acta polytechnica ctu proceedings doi:10.14311/app.2016.3.0039 acta polytechnica ctu proceedings 3:39–42, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app database finliv – focus on staircase method maxim lutovinov∗, jan papuga, milan růžička faculty of mechanical engineering of czech technical university in prague, technická 4, prague, czech republic ∗ corresponding author: maxim.lutovinov@fs.cvut.cz abstract. this paper introduces finliv database focused on gathering, manipulating and providing experimental static and fatigue data. one of its new features is the ability to include also fatigue tests realized by the staircase method. keywords: fatigue, database, staircase method, fatigue limit. 1. introduction finliv is a database intended to gather the information about static and stressand strain-controlled fatigue experiments and also about parameters of various static and fatigue material models derived from such data. it was developed to resolve the problems of researchers in a lack of quality data, their frequent second-hand use, and the way the data items are selected for various benchmark tests. by now, finliv has a large set of tools for entering and processing data. one of the newest tools is the staircase method [1] that automatically evaluates data from stress-controlled experiments and provides the fatigue limit as a result. 2. finliv finliv is represented by two parts, which closely cooperate, but which also can be used as standalone applications. the first part serves for entering data, while the second part serves for listing them. the second part is implemented as a web database and can be found at www.fadoff.cz/page/finliv. it allows the user to have permanent access to the raw experimental data or to the final regression curves retrieved from them. in the web part there are no tools to process the data, it simply retrieves and lists them on a request in a raw or processed form. the first part, finliv.vba [2], was built as an ms excel application written in visual basic for applications, and its main purpose is to enter experimental data and to process them. data processing is necessary in order to e.g. perform regression analysis or to prepare automated input for fatigue solvers, if a particular experimental data set has to be used within a benchmark of a particular calculation method. the ms excel platform was chosen because of its prevalence among engineers, and because it is substantially cheaper than matlab. the application can be accessed from any computer which has ms excel installed. the main interface of finliv.vba is represented by "form" sheet (see the part of the input interface in figure 1). it consists of a group of tables, combo boxes and buttons. data input is realized by filling in the salmon colored cells and by choosing options from the combo boxes. in the first non-empty column there are designations that correspond to the input parameters. the same designations are used in the datasheets (figure 2), where the material data are stored. the datasheets are placed in the same workbook as the “form” sheet, and their structure is quite simple: in the second column there are designations that represent the entered parameters, while the other columns contain information on individual curves of the particular data set. finliv works with data from static experiments and stressand strain-controlled fatigue experiments. since each type of experiments uses a different set of parameters, the “form” sheet changes its structure to accommodate the selected type. in addition to entering data, finliv.vba can also process data. with the help of the application the user can easily perform regression analyses of experimental data and obtain corresponding regression parameters. for strain-life data there are two types of a regression analysis. the first one is carried out in a conventional way, which is represented by linear regression using the least square method. the second method performs non-linear regression according to the 3-d method [3]. the 3-d method is based on the assumption that one value of the total strain amplitude corresponds to only one stress amplitude and to a single number of cycles. if this assumption is valid a curve can be drawn in a 3-dimensional space which besides formulation of the total characteristics depicts also plastic and elastic strain. unfortunately, implementing such a non-linear 3-d regression in excel is problematic, therefore it was decided to carry out the regression analyses in an external matlab application. to prepare an input file for that application the user should click "input for regression" button. to load the resulting parameters of the regression analysis to finliv, the user should click "get regression data" button and choose the output file created by the matlab application. for stress-life data, there are also two types of a regression analysis, but unlike the case of strain-life 39 http://dx.doi.org/10.14311/app.2016.3.0039 http://ojs.cvut.cz/ojs/index.php/app m. lutovinov, j. papuga, m. růžička acta polytechnica ctu proceedings figure 1. "form" sheet of finliv with sample data. figure 2. a datasheet of finliv with sample data. data both types of the regression are carried out in the excel environment. regression analyses of s-n data as well as a visual display of obtained e-n curves are available in another workbook that is generated automatically from finliv. individual sheets in that workbook correspond to the data sets that were selected for processing. at each page of such data there are two plots of regression curves. in the case of strain-life data, there are two manson-coffin curves, constructed on the basis of parameters obtained from conventional regression analyses and 3-d analyses. in the case of stress-life data, the first regression curve is represented by the basquin relation and another one by kohout-vechet relation. for s-n data, a new tool is also available in the same workbook the staircase method, which is described below. 2.1. staircase method the staircase method is a standardized method [1] for carrying out and evaluating experiments in order to determine the fatigue limit related to a given number of cycles. the essence of the method is to increase or decrease the load amplitude of a new test specimen by a constant pre-selected value based on the response of the previous specimen, i.e. whether a failure occurred during the previous experiment. for evaluation, the staircase method uses a special table (figure 3) constructed according to the standard iso 3800 [1]. the rows of the table represent stress levels and the columns represent the order of the experiments. a symbol “o” is entered into the appropriate cell when a failure doesn’t occur, and a symbol “x” in the opposite case. if the essence of the method has been respected, then the symbols “x” and “o” are arranged in the table into the sequence that 40 vol. 3/2016 database finliv – focus on staircase method figure 3. process of generating the output table according to the staircase method in finliv. looks like stairs. in case of a suitable choice of the first stress level and a step size between individual levels, the stress levels after several experiments (about 10 experiments) will oscillate around a certain value – the fatigue limit. the fatigue limit and the corresponding standard deviation can be then calculated with the help of the equations given by the standard [1], which uses values from the left part of the table (figure 3, columns 3 to 8). the fatigue limit labeled as sig_50 in figure 3 is calculated according to: sig_50 = σao + ∆σall( a c + x) (1) here, σao is the lowest load amplitude level at which the first non-failed specimen appears, given that the total sum of the non-failed specimens is less than the total sum of the failed specimens. in the opposite case, the load level corresponding to the first failed specimen should be input. the standard doesn’t specify how to proceed in the case when the total sums of the failed and non-failed specimens are equal. the observations show that in such case the result will be the same regardless of whether failed or non-failed specimens are selected as the reference type. the algorithm implemented in finliv in cases like that calculates the fatigue limit in two steps, with the failed and non-failed specimens as the reference type and then chooses the lowest fatigue limit. the parameter ∆σall is the step size between individual load levels. a and c are sums of the values in the columns z*f and f. the parameter x is equal to +0.5 when the number of the non-failed specimens is lower than the number of the failed specimens, or -0,5 in the opposite case. it is worth noting that the method can be applied to the load levels represented by stress as well as to the levels represented by forces. the result is then the fatigue limit or the force corresponding to the fatigue limit. the parameter s(sig) is a standard deviation of the distribution and is calculated according to: s(sig) = 1.62∆σall( ce − a2 c2 + 0.029) (2) where e is the sum of values in the column zˆ2*f. the staircase method was extended in finliv, and the application analyzes whether the data are suitable for being processed by the staircase method at all. thanks to that, the method can also be used on data that were obtained from experiments initially unintended to be processed by the staircase method. if the data are suitable for the method, finliv automatically draws an output table (figure 3). the algorithm that analyzes data from the selected load levels assembles various combinations that differ 41 m. lutovinov, j. papuga, m. růžička acta polytechnica ctu proceedings in the step size between levels and/or by the number of levels. then individual combinations are checked according to the criteria that say whether individual sets of load levels and corresponding sets of specimen states (failure, non-failure) are suitable for evaluation by the staircase method. these criteria are derived from the essence of the method. for example, one of the criteria says that if the number of the non-failed specimens at the highest load level is more than one, then such set of load levels and specimen states is not suitable for the staircase method. according to the method, the appearance of a non-failed specimen leads to increasing the value of load amplitude. if the experiment with such specimen is not the last one, then the next experiment should be carried out at a higher level of load amplitude, which means that the level of load amplitude corresponding to the non-failed specimen in question is not the highest level in the particular set of experiments. in the other case, if the experiment at the highest load level with non-failed is the last one, then at the same level there cannot be other non-failed specimens. after selected data are checked by the criteria, finliv provides the user the list of possible combinations of load levels and corresponding steps to choose from (figure 3). the combinations are ordered by the number of included specimens. for each combination there are also listed corresponding step sizes and calculated fatigue limits. while selecting a combination it is important to check the ratio of the fatigue limit and the step size. the smaller is the ratio, the more accurate the estimate should be. the drawback of the implemented check is that it works only with load levels and can’t take into account the possibility of individual specimens being removed from evaluation in order to assemble a suitable combination of the load levels and specimen states. 3. summary the paper refers to finliv database for storing, listing and manipulating experimental static and fatigue data. it describes its newest feature, which is the ability to process data by the staircase method. acknowledgements the authors thank for the support of their research by sgs14/181/ohk2/3t/12 ctu grant. references [1] iso 3800:1993(e). threaded fasteners axial load fatigue testing test methods and evaluation of results. switzerland: international organization for standardization, 1993. [2] j. papuga, m. lutovinov. help for finliv.vba excel database [ver. c], fme ctu in prague & evektor, spol. s.r.o., prague 2014. [3] a. nieslony, c. dsoki, h. kaufmann, p. krug. new method for evaluation of the manson–coffin–basquin and ramberg–osgood equations with respect to compatibility. international journal of fatigue 30(1011):1967–1977, 2008. doi:10.1016/j.ijfatigue.2008.01.012. 42 http://dx.doi.org/10.1016/j.ijfatigue.2008.01.012 acta polytechnica ctu proceedings 3:39–42, 2016 1 introduction 2 finliv 2.1 staircase method 3 summary acknowledgements references 323 acta polytechnica ctu proceedings 2(1): 323–324, 2015 323 doi: 10.14311/app.2015.02.0323 concluding remarks d. v. bisikalo1 1institute of astronomy ras, moscow, russia corresponding author: bisikalo@inasan.ru keywords: cataclysmic variables. there are four persons making the concluding remarks of the conference, so i restricted myself by the 1-st part when we discussed the cvs. i am pleased to note that the scientific level of our meeting has been quite high. i have heard all 46 talks on cvs and only one has not been new for me, my own ”magnetic cvs behaviour: a review”. i would like to stress the very good organization of the conference program, which, in particular, appeared as a balanced number of reviews and contributed talks. i was really impressed by the review talks made by franco giovannelli & lola sabau-graziati, hans ritter & ulrich kolb, joseph patterson, margaretha pretorius, pietro parisi, christian knigge, edward sion, paula szkody, klaus reinsch, and solen balman. all these talks contain the absolutely comprehensive material available today on the discussed problems. i do not reproduce their content, since all of them can be found in this book of proceedings, but i want to note that any researcher, studying cvs, independently on the age and scientific status, will find them useful. nonetheless i would like to focus on one of the important problems addressed in the talks. this is the problem raised by hans ritter. we all know that during the last 30 years the rkcat (undoubtedly, the most known catalogue of cvs and related objects) has been annually supplemented at a rate of approximately 9% a year. however, in his talk, hans ritter has drawn our attention to two important points: 1) servicing the data base is now practically a full-time job, 2) he will not be able to provide this service any longer. at this point a natural question for the entire binary community arises, who will continue with rkcat and in what form?”. happily the authors of this catalogue still intensively work and we have time to solve this problem. but, of course, we all should think about it. the contributed talks have been well prepared and delivered, and the reader may see it when reading the papers in the book of proceedings. like for the reviews i will not retell the content of even the most interesting talks, but i will try to pick out general tendencies that came out after analysis of the contributions. besides, i will highlight the results that might be useful for a broader community of researchers, studying objects and problems other than cvs. the first point, i would like to mention, is grand astronomical progress in south africa. its scientists have presented 7 talks, some of which delivered by very young persons, and all of them have excited an intense interest of the audience. undoubtedly, this burst of activity is concerned with the development and launch of the salt. and, in spite of the fact that david buckley still regards the obtained results as an appetizer, they obviously have become the main course in various fields of astrophysics. the presented talks revealed that theoretical studies and interpretations (both in the number of delivered talks and excited interest) noticeably dominate over observations. indeed we have heard wonderful analytical and numerical contributions that not only explain available observational data both from the space and ground-based instruments but also possess the great potential for the predictions, analysis and generalization of upcoming data. i do not want to focus on a particular work since i hope that the reader will first get familiar with the original papers but not these concluding remarks however, i must emphasize noticeable progress in observational methods. this has become my discovery that by using quite simple modifications one can significantly improve the quality of doppler tomography (see the papers of e. kotze, m. uemura, s. zharikov, s. potter, d. kononov) observational works have mostly followed a conservative way of data accumulation, which, undoubtedly, is the most important part of any research and the basics of scientific progress. at the same time we have got some surprises: .odendaal has told us that cal83 is a new candidate to ae aqr-like system; j. v. hernandez santisteban et al. have confirmed the detection of a sub-stellar donor star in a cv; d. chochol et al. have reported the superoutburst observations of a possible new helium-rich dwarf nova 2013 in hercules. the next important aspect in the development of observations is concerned with the effective usage of surveys (see, e.g., the wonderful paper of p. szkody about cvs from sdss and the paper of p. cartera about cvns in 323 http://dx.doi.org/10.14311/app.2015.02.0323 d. v. bisikalo sdss) and data from space telescopes (see e.g. the papers of p. parisi, s. balman, m. kotze, v. suleimanov, a. semena). resuming my concluding remarks i would like to stress that the conference was not only interesting, but also useful. i thank the organizers of the conference not only for the nice time, spent in the marvelous place, but also for the opportunity to discuss all the results in the friendly atmosphere. as i found when talking with many participants, we all are interested in the continuation of this series of meetings and in making this conference recurrent. franco giovannelli & lola sabau-graziati and their team are the main engine of the conference and i hope that they will continue this hard job, the organization of cvs conferences. 324 326 acta polytechnica ctu proceedings 1(1): 326–328, 2014 326 doi: 10.14311/app.2014.01.0326 some personal conclusions sergio colafrancesco1 1university of the witwatersrand, johannesburg (south africa) corresponding author: sergio colafrancesco. email: sergio.colafrancesco@wits.ac.za abstract i present here some personal considerations on the main theoretical, observational and technological challenges offered to the discussion during the meeting ”multifrequency behaviour of high energy cosmic sources”. 1 the theme this meeting presented a wide collection of excellent results on a wide range of subjects in modern multifrequency astrophysics. i am not able to remember a single talk that did not produce a vibrant discussion in and out of the audience. beyond any attempt to summarize the outcomes of the meeting, it seems more appropriate for me to briefly present here some personal considerations on the main challenges offered to the discussion during this meeting. 2 the challenges one of the main aspects of this meeting is its capacity to induce global discussions on some of the most important challenges in astrophysics and cosmology. these challenges address the nature of fundamental questions in the physics of the universe, the frontier of exa-scale data mining and analysis and the technological challenges related to the construction of the largest astronomical facilities in the next decade. 2.1 theoretical challenges: fundamental questions big theoretical challenges regard two important aspects of the structure and evolution of our universe: the cosmic origins and the cosmic extremes. origins. beyond the successes of the standard cosmological scenario, we are still facing some crucial questions: what happened at the beginning of the universe? inflation and precision cosmology. there’s been incredible progress recently in finding the traces that inflation left behind and upcoming experiments on cmb polarization promise to provide even more evidence of what happened during the universe’s infancy. through its sensitivity to gravitational waves, the study of the cmb provides a glimpse into the state of the universe just ∼ 10−35 seconds after the beginning and of physics on grand-unification-theory (gut) energy scales around ∼ 1016 gev, some ≈ 13 orders of magnitude above the energies achievable by current terrestrial particle accelerators. a gravitational-wave background in the early universe would leave a unique, oddparity pattern of polarization in the cmb (b-modes), the magnitude of which is characterized by the tensorto-scalar ratio r. a gwb is generically predicted to exist by inflationary theories, and the current generation of cmb polarization experiments will probe the interesting parameter space of r ∼< 0.05 corresponding to single-field inflationary models at gut scales. the challenges here seem to be a statistically confident measurements of b-modes and the imperative study of the systematics that are present in such a difficult measurement. a multi-frequency study of b-modes with extremely high sensitivity, over a wide range of frequencies and large sky areas would likely be able to measure in the not-so-distant future the value of r with a ∼ 10% precision and will open the field to a deep theoretical exploration of inflationary cosmologies. what is the dark matter? this is a particularly exciting time for dark matter study because there are some intriguing clues pointing to where dark matter particles might be hiding. these clues are helping researchers develop a variety of searches. the three major strategies are direct detection, collider production and indirect detection. while there is a mounting scientific frustration in the non conclusive evidence for the detection of dm particles with direct and collider detection techniques, the sensitivity and the specific strategy of the next coming astronomical observatories (mainly the cta and the ska) will probably allow in the next decade to shed a conclusive word on the nature of dm, or close a large fraction of the dm parameter space and thus open the way to a more detailed exploration of realistic alternatives of the theory of gravity. 326 http://dx.doi.org/10.14311/app.2014.01.0326 some personal conclusions what makes up the rest of the universe? dark energy. in the past 15 years or so, scientists have realized that the ”stuff” making up all the atoms in all the galaxies, stars, planets, and humans we have ever observed only constitutes about 5% of the universe. while we might be close to pinning down the nature of part (≈ 27%) of the missing stuff (e.g., dark matter), what we know about the dominant (≈ 68%) component of the universe, (named historically dark energy for a vague analogy with dm) is still almost nothing. observational efforts are planned (e.g., euclid, ska, des, lsst, etc.) and theoretical exercises to narrow down the available de parameter space are the subject of a restless activity, but a detailed physical characterization of the ”dark energy” is still missing and probably this theoretical activity will be very relevant in pointing at the true nature of the physical mechanism providing the late-stage cosmic acceleration. extremes. where did that come from? cosmic rays and intergalactic particle accelerators. after a century of study, researchers still struggle to understand the origin of cosmic rays, and especially those of ultra high energy. we believe these extreme high-energy oddities play a key role influencing the physics and chemistry that form stars and planets, and even influence life on earth by occasionally causing mutations in dna. and yet, the exact ways in which cosmic rays are accelerated remains a major open question. we’ve discovered where many come from within our galaxy, but the most extreme cosmic rays continue to confound us. the need of a very large collecting area experiment for the extremely rare uhecr events poses several observational challenges, but its delivery will probably open the way to a physical clarification of their origin and of the production sites. challenges to particle acceleration theories in cosmic structures are stronger and stronger and the scientific qualification of the possible product sites is reducing more and more towards regions of compact objects with extremely strong gravitational fields and magnetic fields. what can compact objects teach us? black holes and neutron stars, extreme physics in small packages. enormously powerful gravitational fields that warp the local fabric of space and time. incomparably strong magnetic fields that can stretch atoms themselves into long spindles. materials so dense a teaspoonful would weigh billions of tons. these are just some of the exotic properties of compact objects, a catch-all term for several types of unbelievably dense and remarkable objects, like white dwarfs, neutron stars and black holes. compact objects are known to possess some of the most extreme physical properties ever observed. scattered throughout our galaxy and agns, these objects serve as astrophysical laboratories that test the very limits of physics as we know it. 2.2 observational: big data science it is now becoming clear that the answers to big questions in astrophysics and cosmology requires to address the challenge of dealing with big data quantities, analysis and transport. how will we make sense of it all? astronomically big data. astrophysics and cosmology deal with big everything: big datasets, big simulations and big collaborations. we have information on billions of astronomical objects, and expect to make measurements of many billions more in the next decade. yet the challenge with such a large dataset is not so much in handling its size, but in its complexity. the struggle in trying to find a single rare star in a haystack of billions of nearidentical stars, or understanding the relationships between every single galaxy in the universe, goes beyond simply the enormous number of gigabytes. as more and more data piles up, the teams who are most innovative about analyzing and combining those datasets will be the ones who will likely make the biggest discoveries. instruments like the ska will fully live into the era of big data complexity and will have to address not just the complex techniques of data reduction and analysis, but also the challenges of data distribution and transport over inter-continental distances and yet with high computing capacity on our everyday desktop computer. the next generation of astronomical surveys (and even single observations) will contain information that spans multiple frequencies, and many epochs in time. however there is not yet a next generation of catalog than can describe these surveys. a more meaningful description of catalog data will ultimately lead to better comparisons of astronomical surveys and more robust scientific output. it is not unreasonable that a stratification of various computational -tasks and a possible -re-definition -of -the -data -analysis -objectives and -techniques -will -be -(re-)considered in the light of the experimental and theoretical challenges of the next decade big astronomy science. in this context, small or mid-size multifrequency data centers will probably play a major role in developing innovative applications for (astronomical) data science and producing cutting edge scientific exploitation of the data achieved by large astronomical facilities. 2.3 experimental: big facilities the future large astronomical projects will live inherently in the era of multifrequency astronomy. examples of this synergy are provided by the ska, the cta, the lsst, and the largest future space missions like e.g. 327 sergio colafrancesco euclid, millimetron. the wealth of planed astronomy projects is such that it seems that the future of astronomy can only be limited by the financial constraints of international funding available. both ground-based and space-borne experiments are facing challenges specific to their different nature. some of these are technical and will give rise to extensive technological developments and innovations. some other are programmatic and are related to the complex and ambitious nature of future projects. it is not unlikely that the complementarity between large-scale projects and pathfinders will provide the scientific community the clarity on the technological routes to be followed for the construction of the most successful and challenging astronomical facilities of the next decade, as well as the directions in the optimal integration of these facilities in the most successful multifrequency strategy. this seems to be mandatory in order to answer the remaining big questions on the nature and the evolution of our universe. 3 predictions based on the decennial successes of this meeting, as also testified by this specific one, it is not difficult to predict that the next meeting of this series will provide an even more extended discussion on the major challenges of multifrequency astrophysics, cosmology and astro-particle physics in the next future. acknowledgments s.c. acknowledges support by the south african research chairs initiative of the department of science and technology and national research foundation and by the square kilometre array (ska). 328 the theme the challenges theoretical challenges: fundamental questions observational: big data science experimental: big facilities predictions 261 acta polytechnica ctu proceedings 2(1): 261–263, 2015 261 doi: 10.14311/app.2015.02.0261 properties of recurrent nova t pyxidis based on 2011 outburst k. tanabe1 1faculty of bioshere-geoshere science,okayama university of science,okayama 700-0005 japan corresponding author: tanabe@big.ous.ac.jp abstract we reexamine the properties of the recurrent nova t pyxidis based on our own spectroscopic data accompanying with the photometric ones by vsolj (variable star observers league in japan) during 2011 outburst. one of the purpose of this paper is whether a missing outburst could be happen around 1988-1989. comparing the 2011 outburst data with previous ones, we may conclude that any essential difference can not be found. accordingly it is difficult to deny a small possibility of a ”missing” outburst from 1988 to 1989, taking into account the seasonal gap in its observation for northern hemisphere observers . the problem whether im normae belongs to be a member of t pyx subclass or not is to be postponed by its next outburst taking into account of t pyx’s peculiar spectral behavior. keywords: cataclysmic variables recurrnt novae t pyxidis optical spectroscopy photometry. 1 introduction recurrent novae (rne hereafter), whose definite total number is just 10 (see warner (2008) for example), are divided into three distinct subclass, namely t pyx subclass, u sco subclass and t crb subclass. among these three, t pyx subclass has only one member, t pyxdis itself and has a quite unique property of very short orbital period (1.8 h). t pyxidis is the only galactic nova belonging to the southern constellation pyxis. outburst of this object was discovered by h. leavitt in 1902 photographically. this rn has experienced 5 outbursts (1902, 1920, 1944, 1966) including 1890 detected on the harvard archival plate, before 2011 outburst. while detailed photometric observations were performed before 2011 outburst (see schaefer (2010)), spectroscopic observations performed from early time were rather fragmentary. moreover this rn accompanies nebulosity. the 2011 outburst of the rn t pyxidis was detected after 45 year of absence (twice as long as recurrence period) by linnolt on april 14.2931(ut), 2011(waagen(2011)). im mediately after discovery we started low-resolution spectroscopy and obtained detailed feature of this rn before maximum light. the obtained results is reported in the paper by imamura and tanabe (2012). on the other hand the results of the photometric observations are based on the results performed by the members of vsolj (variable star observers league in japan). in section 2, we summarize the spectroscopic observational results by imamura and tanabe (the present author), including both the basic properties of t pyx obtained before 2011 outburst and the essentials of socalled tololo classification scheme. in section 3, we try to compare the photometric observations by vsolj during 2011 outburst with the template light curve by schaefer (2010) based on its previous outbursts. then we discuss a possibility of ”missing outburst” to be expected in 1988-1989 by comparing 2011 outburst with the previous ones by using the schaefer’s template light curve (schaefer (2010)). also we mention about the problem whether im normae can be another member of t pyx subclass. at the end of this section we summarize our results. 2 summary of low-resolution optical spectroscopy 2.1 basic properties of t pyxidis based on pre-2011 outburst as it is mentioned above, t pyx is a quite unique rn for its very short orbital period (1.83 hour) compared with other rne. photometric behavior is characterized by slow rise and slower decline. its position is (α,δ)=(9h04m41.47,-32deg22.0min) and (λ,β)=(150deg,-46deg). this latter ecliptic coordinate makes its seasonal gap (almost half a year) the for northern hemisphere observers. the observed spread of magnitude is b=6.4-15.4. in addition the inclination is about 10 degree (for example, utas et al (2010)). the distance had not been known; recent value by sokoloski(2013) is 4.8±0.5kpc. 261 http://dx.doi.org/10.14311/app.2015.02.0261 k. tanabe 2.2 spectral classification of novae modern and phenomenological spectral classification of classical novae(we denote cne hereafter) is established by williams (so-called tololo classification system). according to this system, spectra of cne are divided into two classes, namely fe ii class and he/n class. the former corresponds to slow evolution novae and the latter to fast ones. as this phenomenological scheme is thought to be a reflection of the physical stage of nova ejecta, it seems to be applicable not only to cne but also to rne. 2.3 spectral evolution of t pyxidis in its early stage of 2011 outburst we started our low-resolution (r∼400) spectroscopy, using dss-7 spectrograph with st-402 ccd camera (both of them of sbig (santa barbara instrumental group)) production attached to a 28 cm schmidt cassegrain telescope (celestron production) at tanabe’s personal observatory(tpo) located in the north west edge of okayama city in japan (many photometric data obtained at tpo on dwarf novae are published by vsnet collaboration team). this site is quite suitable for astronomical observation because of its fine weather and good seeing. the okayama astrophysical observatory (oao) of national observatory in japan (naoj) is located at 40 km west from tpo. figure 1: spectral evolution of t pyxidis. upper panel denotes earlier stage,which shows he/n feature. on the contrary lower panel is later stage that shows typical feii spectral phase. the date after maximum light (minus means before maximum) are denoted on the right shoulder on each spectrum. spectral observations started at two days after discovery (april 16th and ended at 14th of may due to its position close to the sun. total observation are 13 nights, among which 9 nights are before the maximum light. details of the results are reported in imamura and tanabe (2011). here we only summarize the peculiarity about the pre-maximum behavior that the spectral evolution is not only hybrid but opposite direction from he/n phase to fe ii one. this feature is quite exotic (see figure 1). according to the obtained p cygniprofle, the expansion velocity turns from decreasing to increasing. such a transition seems to cause this extraordinary evolution from he/n to fe ii phase. the relation between this evolutional feature and the temporal change of physical state as a binary system will be discussed in a separate paper. 3 discussion 3.1 ”missing” outburst? as is shown in figure 2, the photometric behavior of t pyx during its 2011 outburst seems to be essentially the same as the template curve combined by previous outbursts (schaefer (2010)). taking into account that 45 year interval is twice as averaged interval from those 5 outburst records, we feel temptation to find out evidence of outburst around 1988-1989. in addition we can find no incompatible feature on our spectra with past fragmentary spectral data. hence it is plausible that the structure of the t pyx system does not change in binary structure after 45 years absence from outburst. therefore it is worth seeking the possibility of ”missing” outburst around 1988-1899. we tried to search the photometric data in vsolj database and found several records from the end of 1988 to 1989 with visual magnitude of 13 mag. however, unfortunately, these records were by a single observer and no distinct light variation could be seen. so this cannot be a strong evidence for missing outburst. figure 2: light curve obtained by vsolj members during 2011 outburst. smooth curve denotes a template light curve by schaefer (2010) based on previous photometric (b magnitude) data. we can see coincidence between two and conclude that no essential difference exist. 262 properties of recurrent nova t pyxidis based on 2011 outburst 3.2 is im normae a real member of t pyx subclass? recurrent nova im normae is a possible member of t pyx subclass from the point of view of light curve and its orbital period (schaefer (2010)). however taking into account of the spectroscopic properties based on 2011 outburst, spectral evolution of t pyx subclass is to posses such a property as the inverse change of normal transition from fe ii phase to he/n one. to confirm that im nor really belongs to t pyx subclass, it is necessary for us to wait the next outburst (some 80 years) for obtaining a new data on its spectral evolution. 3.3 conclusions 1) during the early stage of outburst, t pyx shows the hybrid spectral transition from he/n phase to fe ii. 2) light curve obtained by vsolj which seems to be identical to the schaefer’s template curve suggests the missing 1988-1989 outburst. but no strong evidence exists during these years. 3) without spectroscopic data, the problem whether im normae really a member of t pyx subclass is not settled. acknowledgement the author thanks vsolj members (leader seiichiro kiyota) for permitting us to make use of photometric data. minako ogi, my graduate student, helped me for her assistance of making a figure 2. the author also thanks naoko tanabe for her preparation of the instruments during our spectroscopic observations. references [1] leavitt,h.1913,harvard college obs.cir.179,1 [2] imamura,k and tanabe,k.: 2012 pasj 64,l9 [3] schaefer,b. : 2010, apjs, 187, 275 doi:10.1088/0067-0049/187/2/275 [4] utas et al. 2010,mnras, 409,237 doi:10.1111/j.1365-2966.2010.17046.x [5] sokoloski et al. 2013,apjl 770,l33 doi:10.1088/2041-8205/770/2/l33 [6] waagen,e.et al. 2011,cent.bur.electron.telegram,2700 [7] warner, b. 1995, in cataclysmic variable stars (new york: cambridge university press) doi:10.1017/cbo9780511586491 [8] warner, b. 2008, in classical novae, ed. bode, m. f. & evans, a. (new york: cambridge university press) [9] williams, r. e. 1992, aj.104, 725 [10] williams, r. e. 2012,aj 144, 98 discussion ashley pagnotta: there was no missed eruption in the year of 1980s.too many observers were working.brad schaefer has worked at all possibble archival amateuer’s data looking for trails hasn’t forward anything. kenji tanabe: i asked recently the observer t.kato for his 1988-1989 vsolj data whether the value of 13 magnitudes data indicating this rn’s high state is reliable or not.his answer was of little confidence. he said that the telescope used was 20cm and visual observation at his early day of vs observations. at first i thought this was performed using 60cm telescope of kyoto university. 263 http://dx.doi.org/10.1088/0067-0049/187/2/275 http://dx.doi.org/10.1111/j.1365-2966.2010.17046.x http://dx.doi.org/10.1088/2041-8205/770/2/l33 http://dx.doi.org/10.1017/cbo9780511586491 introduction summary of low-resolution optical spectroscopy basic properties of t pyxidis based on pre-2011 outburst spectral classification of novae spectral evolution of t pyxidis in its early stage of 2011 outburst discussion "missing" outburst? is im normae a real member of t pyx subclass? conclusions acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0022 acta polytechnica ctu proceedings 4:22–26, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app thermal-hydraulic analysis of irt-4m in reactivity insertion accident at vr-1 reactor filip fejt faculty of nuclear sciences and physical engineering, czech technical university in prague, czech republic correspondence: filip.fejt@fjfi.cvut.cz abstract. the paper deals with thermal-hydraulic analysis during reactivity insertion accident, i.e. a step increase of nuclear system reactivity by 0.7βeff, at vr-1 reactor. the reactor utilizes irt-4m type of fuel assemblies, and even though these fuel assemblies are designed for an operation at the high-power research reactors, they might be also used for zero-power reactors. the thermal-hydraulic analyses must take into account several specific assumptions that are derived from vr-1 reactor specifications. the reactor does not require a forced water flow for a fuel cooling, the core is placed in an open vessel with atmospheric pressure, and amount of coolant water in the vessel is sufficient for providing the inlet water at room temperature for the whole event. coolant circulation is expected to be formed only by natural convection. keywords: irt-4m, ria, relap5, trace5, natural convection. 1. introduction thermal-hydraulic analysis is a necessary part of every safety analyses report. the adequate cooling of the fuel assemblies must be ensured to maintain the core integrity during normal and abnormal operational transients, and accidents. in general, the research reactors may be designed for both natural convection and forced convection. the pool type research reactor vr-1 uses a natural convection of coolant to remove a core heat. the upward flow in natural convection is caused by different fluid densities of heated and unheated volumes. important point of safety analyses is an onset of nucleate boiling (onb) [1] that is defined as a difference between the fuel surface temperature at the starting point of a nucleate boiling and the fuel surface temperature at a local cooling condition. reaching a nucleate boiling significantly limits the heat transfer leading to a dramatic increase of the fuel surface temperatures in order to maintain the total heat flux. the main purpose of thermal-hydraulic analysis is to inspect all of possible events that may lead to an accident, and verify that normal and abnormal operation has safety margins before the critical level is reached. 2. thermal-hydraulic codes due to the various research reactor designs, i.e. operational mode, fuel type and applications, each reactor has a unique thermal-hydraulic characteristic. computational codes must be capable of dealing with a broad range of different designs. since the current codes are still based on dimensionless quantities that are used to calculate flow patterns in different fluid situation, the equations are mostly derived from simplified benchmarks and loops that must be considered similar to a problem of interest. although many well-known codes, e.g. relap, trace, paret, have been designed for power reactors at first, they have been developed into general calculation tools that may be used even for research reactors. the main phenomena of vr-1 reactor analysis is natural convection, and therefore the chapter is followed by a brief description of natural convection part of the codes. 2.1. relap5 the relap5/mod3.3 code has been developed for best-estimate transient simulation of light water reactor coolant systems during postulated accidents [2]. the code includes several behaviors effecting a reactor coolant, e.g. loss-of-coolant accident, anticipated transient without scram, loss of feedwater, loss of flow etc. unlike its predecessors, relap5 series uses a two-fluid, nonequilibrium, nonhomogeneous, hydrodynamic model for transient simulation of the two-phase system behavior. convection calculations rely on evaluating forced convection, laminar convection, and natural convection and selecting the maximum of these values. the correlations are by dittus-boelter [3], kays [4], and churchill-chu [5] respectively. 2.2. trace trace has been designed to calculate best-estimate analyses of loss-of-coolant accidents, operational transients, accidents in pressurized light-water reactors and boiling light-water reactors [6]. the code includes multidimensional two-phase flow, nonequilibrium thermo-dynamics, generalized heat transfer, level tracking, and reactor kinetics. the partial differential equations that describe twophase flow model are solved using finite volume numerical methods. components are solved in one-dimension 22 http://dx.doi.org/10.14311/ap.2016.4.0022 http://ojs.cvut.cz/ojs/index.php/app vol. 4/2016 thermal-hydraulic analysis of irt-4m in ria at vr-1 or three-dimensions when necessary. timesteps are limited by courant-limit factor to ensure the computational stability of an examined system. the courantlimit depends on a phenomena and current conditions in a system. description of this factor can be found in [2]. similar to relap5, the trace follows a component-based approach to modeling a reactor system. each part of the system is inserted as a unique component. list of components in the code is designed to cover all power plant models. after that all the components are connected as the user requires. expected mode of calculation in examined ria is pre-critical heat flux region. the correlations are provided for the laminar and turbulent forced convection regimes and for natural convection. the wall heat transfer coefficient for single-phase liquid [6] convection is taken as the maximum of the values for laminar and turbulent forced convection and natural convection (1) hwl = max{hlam,hturb,hnc}, (1) where hnc is the wall heat transfer coefficient for natural convection, hlam is the wall heat transfer coefficient for laminar forced convection, and hturb is the wall heat transfer coefficient for turbulent forced convection. heat transfer coefficient for natural convection uses correlations for both the laminar and turbulent regimes and a maximum of the two values is used to ensure continuity. the heat transfer coefficient is computed from (2) hnc = kl dh · nunc, (2) where k is thermal conductivity evaluated at the bulk fluid temperature and dh is the hydraulic diameter. the nusselt numbers are given by (3) and (4) nunc,turb = 0.1 · (gr l · pr l)1/3, (3) nunc,lam = 0.59 · (gr l · pr l)1/4. (4) both grashof and prandtl numbers are evaluated at bulk fluid temperature. 3. vr-1 reactor model the vr-1 core is placed in a vessel with large water reserve. during the transient the water reserve ensures the room temperature at the core inlet. even though the transient may last several hundred seconds, the average water temperature in the vessel will not increase because the flow rate caused by natural convection is very small. in order to maximize the first power peak, the coolant is expected to be stationary at the beginning of the simulation. this condition can be achieved when the main pump in the vessel is turned off and the reactor is operated at the very low power causing minimal flow velocity. the irt-4m fuel assemblies used the vr-1 reactor consist of 4, 6, or 8 fuel layers arranged in rounded rectangular tubes [7]. each fuel layer is surrounded by cladding layer from both inner and outer side. cladding layers are made of aluminium similar to fuel layers where the uo2 is dispersed also in aluminium. a thin layer thickness, both the fuel and the cladding, in addition to a high thermal conductivity of the aluminium leads to a very flat radial temperature profile in fuel assemblies. the core is represented by an averaged 8-tube fuel assembly (fig. 1) to comply with point kinetic equations that are integrated into the computational codes [8]. figure 1. top view of 8-tube irt-4m fuel assembly (yellow – water coolant, blue – cladding layer, green – fuel layer). 3.1. relap5 specification using the component-based structure of input files, a list of utilized components can be found in tab. 1. the bottom plenum serves as a coolant distributing element at the bottom part of fuel assembly. distributed coolant flows through the fuel assembly (each pipe represents one coolant flow area between heatstructures), and at the top of the fuel assembly the top plenum mixes the heated coolant that flows directly to the vessel water reserves. the model can be seen in fig. 2. component number pipe (fa coolant) 8x pipe (vessel) 1x heatstructure (fuel tubes) 8x plenum 2x table 1. list of relap5 components utilized in the single assembly vr-1 model. 23 filip fejt acta polytechnica ctu proceedings figure 2. relap5 model with green pipes connected to heatstructures, bottom and top plenum, and a large pipe serving as a reactor vessel. 3.2. trace specification unlike relap5 model, the trace model is not using the plenum component for distributing and mixing the coolant because it is assumed to be obsolete in current development. its role has been taken by advanced pipe settings; unfortunately, the low dimensions of fuel assembly flow surfaces and complex flow distribution made it inapplicable in this model. therefore the bottom and top plenum were replaced by independent pipes denying any interaction between coolant coming to/from flow surfaces. similar idea had to be applied also to the pipe representing the vessel; the vessel is divided into 8 pipes. the model can be seen in fig. 3. this pipe separation requires an extended usage of components (tab. 2). component number pipe (fa coolant) 8x pipe (vessel) 8x heatstructure (fuel tubes) 8x pipe (bottom plenum) 4x pipe (top plenum) 4x table 2. list of trace components utilized in the single assembly vr-1 model. 4. results for the sake of clarity, the graphical interpretation of results follows two main rules – relap results are plotted with connected points and -r is added in the legend; trace results are plotted with separate points and -t is added in the legend. each figure 3. trace model with green pipes connected to heatstructures, set of pipes serving as bottom and top plenum, and set of pipes serving as a reactor vessel. spatial value is considered to be averaged over flow surface, therefore such a value has a two digit number representing a number of outer and inner fuel layer (fuel layer number starts from the outside of the fuel assembly: 1, 2, 3, . . . , 8, v denotes central area). 4.1. reactor power output the reactor power and the reactivity of the system can be seen in fig. 4. the reactivity is increased to the value of 0.7βeff at t = 0.5 s, after that the value is decreased by feedback effects created from both the fuel and the moderator. the reactivity increase starting at t = 20 s is caused by the slowly increasing coolant velocity effecting its heat removal capabilities. after 200 seconds the reactivity reaches the zero value. 4.2. coolant velocity fig. 5 shows a fast velocity increase of coolant that causes the reactivity peak shown in fig. 4. this analysis is valid as long as the initial coolant velocity is zero, i.e. vessel pump is turned off, and the reactor power is very low. 4.3. coolant temperatures the temperatures of flow surfaces are shown in fig. 6. no coolant reaches the boiling point during the occurring ria event. it can be clearly seen that both relap5 and trace have a very good agreement in coolant temperatures between outside fuel layers. the difference between the codes is approx. 10 ◦c in the middle of the fuel assembly. 5. conclusions despite the equation differences that can be found in both codes, relap5 and trace code shows a very good agreement in both point kinetics and thermalhydraulic. even though a difference can be observed 24 vol. 4/2016 thermal-hydraulic analysis of irt-4m in ria at vr-1 figure 4. reactor power and reactivity of system during step increase of reactivity accident by 0.7 βeff. figure 5. outlet coolant velocity during step increase of reactivity accident by 0.7 βeff. figure 6. outlet coolant temperature during step increase of reactivity accident by 0.7 βeff. 25 filip fejt acta polytechnica ctu proceedings in the final power, it has no effect on maximal outlet coolant temperature. higher power, i.e. higher heating, is always compensated by an increase of velocity of a coolant that leads to a temperature decrease. the coolant velocity can be compared to negative reactivity feedback effects. the separate pipes can be considered as the most significant difference between both models. this effect can be seen in inner flow cross sections where the temperature difference can reach up to 10 ◦c. since the maximal temperatures are reached in outer flow cross sections and both codes compute the same coolant temperature, it can be safely assumed that the separate pipe model does not effect the maximal temperatures. references [1] j. daeseong, p. suki, p. jonghark, et al. cooling capacity of plate type reserach reactors during the natural convective cooling mode. progress in nuclear energy 56:37–42, 2012. [2] u.s. nuclear regulatory commission. relap5/mod3.3 code manual volume i: code structure, system models, and solution methods, 2010. division of systems research, office of nuclear regulatory research. [3] f. dittus, l. boelter. heat transfer in automobile radiators of the tubular type. publications in engineering 2:443–461, 1930. [4] w. kays. numerical solution for laminar flow heat transfer in circular tubes. american society of mechanical engineers 77:1265–1274, 1955. [5] s. churcill, h. chu. correlating equations for laminar and turbulent free convection from a vertical plate. international journal of heat and mass transfer 18:1323–1329, 1975. [6] u.s. nuclear regulatory commission. trace v5.840 theory manuals: field equations, solution methods, and physical models, 2013. division of systems research, office of nuclear regulatory research. [7] j. rataj, l. sklenka. calculations and measurement at the training reactor vr-2. international atomic energy agency, publications 2007. [8] j. daeseong, p. jonghark, c. heetaek. development of thermal hydraulic and margin anlysis code for steady state forced and natural convective cooling of plate type fuel research reactor. progress in nuclear energy 71:39–51, 2014. 26 acta polytechnica ctu proceedings 4:22–26, 2016 1 introduction 2 thermal-hydraulic codes 2.1 relap5 2.2 trace 3 vr-1 reactor model 3.1 relap5 specification 3.2 trace specification 4 results 4.1 reactor power output 4.2 coolant velocity 4.3 coolant temperatures 5 conclusions references 217 acta polytechnica ctu proceedings 2(1): 217–221, 2015 217 doi: 10.14311/app.2015.02.0217 the slow nova v1280 sco: a short review h. naito1 1graduate school of science, nagoya university, furo-cho, chikusa-ku, nagoya 464-8602 corresponding author: naito@phi.phys.nagoya-u.ac.jp abstract we summarize the results of the extremely slow nova v1280 sco and discuss the approach using discrete multiple blueshifted absorption lines, such as metastable he i* and na i d, detected in our high-resolution spectra. keywords: cataclysmic variables classical novae optical spectroscopy photometry individual: v1280 sco. 1 introduction a nova occurs in a binary system consisting of a white dwarf (wd) and a normal star. the correlation between the wd mass and the decline rate of the light curve is widely accepted (e.g. hachisu and kato 2006), and the nova which occurs on low mass wd tends to decline in brightness at a slow rate and shows slow ionization evolution of its ejacta. a slow nova provides us a good opportunity to perform detailed observation over a long duration. in this review, we summarize the results of the extremely slow nova v1280 sco and discuss the approach using discrete multiple blue-shifted absorption lines, such as metastable he i* and na i d, detected in our high-resolution spectra. 2 summary of v1280 sco 2.1 initial observations v1280 sco was independently discovered by two japanese amateur astronomers (y. nakamura and y. sakurai) at the position of r.a. = 16h57m41s.0, decl. = −32◦20’ 36”.4 (equinox 2000.0) on 2007 february 4 at ninth visual magnitude (yamaoka et al. 2007). it was identified as a classical nova by naito and narusawa (2007) from a low resolution spectrum obtained on february 5.87 (one day after the discovery). it reached the maximum brightness of v = 3.78 mag on february 16, which was 11.3 days after its discovery (munari et al. 2007). the amplitude (a), the magnitude difference between pre-outburst and maximum brightness, is 15 mag or larger because das et al. (2008) noted that no star was visible down to b and r magnitudes of 20.3 and 19.3, respectively, at the position on pre-discovery plates. the notable point is that v1280 sco showed a remarkable formation of dust in its very early phase (das et al. 2007). this dust formation was directly detected using eso’s very large telescope interferometer (vlti) by chesneau et al. (2008) and this result was published as the eso press release (eso0822: http://www.eso.org/public/news/eso0822/). 2.2 infrared observation das et al. (2007) observed the nova in the nir region on 2007 march 4.95, and found that the continuum in the 1.08-2.35 µm region had risen sharply, indicating dust formation in the nova ejecta. das et al. (2008) suggested that the dust was in clumps from nir studies of v1280 sco. puetter et al. (2007) reported spectroscopic observations in the visual-nir regions carried out in 2007 may, and found that the nova was in a very low-excitation state showing strong c i lines and no discernible he i emission. 2.3 high spatial resolution observation high spatial resolution monitoring of the dust formation event was performed using the vlti during the first four months following the discovery, indicating that the dusty shell expanded regularly (chesneau et al. 2008). chesneau et al. (2012) revealed the presence of a dusty hourglass-shaped bipolar nebula around v1280 sco based on mid-infrared imaging observations taken in 2010 and 2011 using the vlt spectrometer and imager for the mid-infrared (visir). 2.4 high quality photometric observation hounsell et al. (2010) published the data set of v1280 sco observed by the solar mass ejection imager (smei) on board the coriolis satellite. thanks to high quality and high time resolution, they revealed that there had been three major but short episodes of brightening near 217 http://dx.doi.org/10.14311/app.2015.02.0217 h. naito maximum light (before 2007 february 20), which had not been detected by any ground-based telescopes. 2.5 distance and wd mass parameters (distance and wd mass) of v1280 published in the literature are controversial (see table 1). hounsell et al. (2010) estimated the distance to be 630 ± 100 pc by measuring the condensation time of dust grains, assuming that the condensation temperature of the dust was 1200 k and the ejection velocity was about 600 km s−1. they also noted that wd mass of 0.6 m� is likely to be an upper limit. on the other hand, chesneau et al. (2008) derived the distance of 1.6 ± 0.4 kpc from direct observations of the size of the expanding shell with velocity of ∼500 km s−1. the discrepancy between these two estimations of distance is over a factor of two, which could result from the complexity in the physical conditions (temperature and velocity) of the dust shell. das et al. (2008) estimated the distance to be 1.25 kpc using the maximum magnitude versus rate of decline (mmrd) relation, and inferred that the higher mass end (1-1.25 m�) may be supported when the amplitude a and the expansion velocity values observed inv1280 sco were taken into account. 3 our results in this section, our results of v1280 sco are summarized. especially we focus on multiple absorption lines detected in high-resolution spectra (see sadakane et al. 2010, naito et al. 2012, naito et al. 2013 for details). photometric and spectroscopic observations have been conducted from february 2007 to july 2013 over table 1: parameters of v1280 sco in the literatures distance [kpc] wd mass [m�] reference 0.63 ± 0.10 < 0.6 hounsell et al. 2010 1.6 ± 0.4 — chesneau et al. 2008 1.25 1.0-1.25 das et al. 2008 1.1 ± 0.5 < ∼0.6 naito et al. 2012 0 2 4 6 8 10 12 14 16 18 0 500 1000 1500 2000 2500 m a g n it u d e s days since maximum v723 cas (vis) －2.8 mag. v1280 sco (v) nebular phase v1280 sco v723 cas figure 1: comparison of light curve and nebular phase between slow novae v1280 sco and v723 cas. light curve of v723 cas is collected from aavso database (http://www.aavso.org) and is shifted by −2.8 mag. v723 cas entered the nebular phase about 18 months after maximum (iijima 2006), while v1280 sco took about 50 months (about three times longer than v723 cas) to enter the nebular phase (naito et al. 2012). 218 the slow nova v1280 sco: a short review six years. photometry in b, v , rc, ic, and y band was carried out with a 0.51-m reflector at osaka kyoiku university. low-resolution spectroscopy (r∼1000) was carried out mainly with 2.0-m nayuta telescope at nishi-harima astronomical observatory and highresolution spectroscopy (r∼60000) was carried out with 8.2-m subaru telescope. according to iijima 2006, v723 cas had been the slowest nova, spending a long time to enter the nebular phase. to compare the evolution between v1280 sco and v723 cas, the light curves and the nebular phases of them are shown in figure 1. v723 cas declines in brightness at very slow rate gradually, while v1280 sco keeps its brightness (v∼10) for 2000 days. v723 cas entered the nebular phase, defined by the appearance of both [o iii] 4959 and 5007, about 18 months after maximum (iijima 2006), while v1280 sco took about 50 months (about three times longer than v723 cas) to enter the nebular phase (naito et al. 2012). considering that v723 cas had the longest transitional time to enter the nebular phase, we conclude that v1280 sco is going through the slowest spectral evolution among known classical novae. hachisu and kato (2004) estimated the mass of v723 cas to be 0.59 m� by fitting their theoretical light curve. our results suggest that the mass of a wd in v1280 sco system might be 0.6 m� or lower by comparing evolution rates between v1280 sco and v723 cas. 0 1 2 3 4 5 5870 5875 5880 5885 5890 5895 5900 in te n s it y + c o n s t. wavelength [ ]å 2009 jun. 16 2010 jul. 1 2011 aug. 6 na d1(is)na d2(is)na d1na d2 2013 jul. 1 he i figure 2: multiple high velocity absorption lines associated with na i d1 and d2. we found discrete multiple blue-shifted na i d lines, ranging from −650 to −900 km s−1, on high-resolution spectra (figure 2). some components weakened significantly from 2009 to 2011 and almost all lines disappeared in 2013. similar absorption lines associated with ca ii h and k and metastable he i* 3188 and 3889 are shown on our spectra. metastable he i* 3889 absorption lines were detected for the first time in 2011 (four years after the maximum light) and have been shown until 2013 (figure 3). this is the first detection of he i* absorption lines in the ejected (circumstellar) gas around novae (naito et al. 2013). we suggest that the complex evolutions of multiple absorption lines are due to combined changes in physical conditions, such as the density, recombination and ionization rate. survival time of these absorption lines in v1280 sco is an order of years, which is much longer than those observed among fast novae (an order of weeks or months; williams and mason 2010). this can be related to the fact that the ionization evolution of v1280 sco is very slow. the behavior associated with he i* in 2011 can be understood that the number of ultraviolet photons had increased significantly to produce singly-ionized helium as the central photosphere shrinks to become hotter, and the disappearance of na i d lines low excitation lines in 2013 can be caused by an additional increase in the number of ultraviolet photons. 0 1 2 3 4 5 3820 3830 3840 3850 3860 3870 3880 3890 3900 in te n s it y + c o n s t. wavelength [ ]å 2009 jun. 16 2011 aug. 6 2012 mar. 20 2013 jul. 1 h e i * 3 8 8 9 h i 3 8 3 5 h i 3 8 3 5 ( h 9 ) h i 3 8 8 9 ( h 8 ) s i ii 3 8 5 6 s i ii 3 8 6 3 f e i i 3 8 2 5 s i ii 3 8 5 4 figure 3: multiple high velocity absorption lines associated with he i* 3889. 4 discussion and conclusions as described in the previous section, metastable he i* absorption lines could survive longer than low excita219 h. naito tion lines, such as na i d and ca ii h and k. this means that metastable he i* absorption lines are more likely to be detected than low excitation lines in the late phase. moreover metastable he i* lines are often used to derive physical parameters of nebulosity because the processes of transition are well understood. these characteristics of metastable he i* lines have advantage in measuring helium composition and the mass of the ejected material, and in studying nova shell structures. using metastable he i* lines in v1280 sco, we show that the ejected shell consists of numerous clumpy gas which cover a significant part of the continuum emitting radiation region as figure 4 (naito et al. 2013). we postulate that this approach is very useful to research nova shell. we reveal that v1280 sco is the extremely slow nova and is available for high-resolution spectroscopic observation for long periods. we attempt to obtain new results of v1280 sco by follow-up observations. free-free emission region clumpy gas photosphere { observer figure 4: schematic of ejected shell producing absorption lines of metastable he i* (see naito et al. 2013 for details). acknowledgement i thank dr. akito tajitsu, prof. kozo sadakane and dr. akira arai for longstanding observations and valuable discussion. i also thank prof. iijima for his useful comments. references [1] chesneau, o., et al.: 2008, a&a 487, 223. [2] chesneau, o., et al.: 2012, a&a 545, a63. [3] das, r. k., et al.: 2007, cbet 866, 1. [4] das, r. k., et al.: 2008, mnras 391, 1874. doi:10.1111/j.1365-2966.2008.13998.x [5] hachisu, i., kato, m.: 2004, apj 612, l57. doi:10.1086/424595 [6] hachisu, i., kato, m.: 2006, apjs 167, 59. doi:10.1086/508063 [7] hounsell, r., et al.: 2010, apj 724, 480. doi:10.1088/0004-637x/724/1/480 [8] iijima, t.: 2006, a&a 451, 563. [9] munari, u., et al.: 2007, cbet 852, 1. [10] naito, h., narusawa, s.: 2007, iau circ. 8803, 2. [11] naito, h., et al.: 2012, a&a 543, a86. [12] naito, h., et al.: 2013, pasj 65, 37. [13] puetter, r. c., et al.: 2007, iau circ. 8845, 1. [14] sadakane, k., et al.: 2010, pasj 62, l5. [15] williams, r., mason, e.: 2010, ap&ss 327, 207. doi:10.1007/s10509-010-0318-x [16] yamaoka, h., et al.: 2007, iau circ. 8803, 1. 220 http://dx.doi.org/10.1111/j.1365-2966.2008.13998.x http://dx.doi.org/10.1086/424595 http://dx.doi.org/10.1086/508063 http://dx.doi.org/10.1088/0004-637x/724/1/480 http://dx.doi.org/10.1007/s10509-010-0318-x the slow nova v1280 sco: a short review discussion alessandro ederoclite: how long after the explosion do you see the emergence of the he i* lines? hiroyuki naito: our latest spectrum taken in july 2013, shows the emergence of the he i* lines which indicates that these lines have appeared until at least 6.5 years after maximum. 221 introduction summary of v1280 sco initial observations infrared observation high spatial resolution observation high quality photometric observation distance and wd mass our results discussion and conclusions 118 acta polytechnica ctu proceedings 1(1): 118–122, 2014 118 doi: 10.14311/app.2014.01.0118 v2282 sgr revisited roberto nesci1, corinne rossi2, antonio frasca3, ettore marilli3, paolo persi1, nicola cornero4 1inaf-iasf, via fosso del cavaliere 100, 00133 roma, italy 2university la sapienza, p.le a. moro 2, 00186 roma, italy 3inaf/oact, catania, italy 4cdso, aosta, italy corresponding author: roberto.nesci@iaps.inaf.it abstract the nature of v2282 sgr is examined on the basis of several multiband observations: a 20 years long i-band light curve of v2282 sgr obtained from archive photographic plates of the asiago and catania observatories; a ccd r-band light curve obtained at cornero observatory; jhk photometry from 2mass and ukdiss; spitzer irac and mips images; optical spectra from loiano observatory; x-ray flux from chandra. the star has a k-type spectrum with strong emission lines and is irregularly variable at all wavebands. the overall evidences suggest that v2282 sgr is a pre main sequence star with an accretion disk. keywords: variable stars pre-main-sequence stellar evolution. 1 introduction the m20 (trifid) nebula is a spectacular star forming region of the milky way: it is projected in the sky over the open star cluster ngc 6514, and its distance has been estimated by several authors, ranging from 1.4 kpc [6] up to 2.7 kpc [1]. the variable v2282 sgr was mainly studied by [5] who reported an amplitude of about 0.5 mag in the i band and classified it as a possible orion type variable. the star is positionally consistent with the soft x-ray source cxom20 180216.8-230347 detected by chandra in m20 [9]; it appeared not to be variable in x-rays, with a flux level of 0.167 c/s, and therefore was not further discussed in that paper. it is also consistent with a spitzer infrared source of the mid-infrared survey of the m20 by [10], classified as a proto-stellar source (class i/0 n.16) without comments about a possible optical counterpart. a nebular variable (either of the t tau, rw aur or t ori type) is believed to be a pre main sequence (pms) star already partially clear of its originating circumstellar envelope [2]: this seems at odds with classification as a class i/0 object, which is a still deeply embedded source. for this reason we decided to explore in more detail its nature. 2 photographic light curve we measured the magnitude of v2282 sgr on 25 plates with i-n emulsion and rg650 filter taken with the 65/90/235 schmidt telescope of the asiago observatory, and on 109 (i-n + rg650) plates of the 40/40/120 schmidt of the catania observatory. plates were digitized with an epson 1680 pro scanner in transparency mode at 1600 dpi. a sequence of comparison stars was defined selected from the gsc2.3.2 catalog, using the n band magnitudes (see fig. 1 and table 1). aperture photometry was made with iraf/apphot. the photometric accuracy of our magnitudes ranges from 0.08 mag at i = 13 to 0.18 mag at i = 14.5: fainter stars have larger dispersion, as expected. the rms deviation of the values of n magnitude for v2282 sgr was 0.25 mag in the asiago and 0.34 mag in the catania datasets, substantially larger than the faintest comparison star. the resulting overall light curve of the asiago and catania observations is shown in fig. 2. v2282 sgr shows substantial variations up to 1 magnitude, with an average level n = 14.0. 118 http://dx.doi.org/10.14311/app.2014.01.0118 v2282 sgr revisited figure 1: finding chart of v2282 sgr from a poss ii plate with iv-n emulsion. the variable is marked with l. comparison stars are labelled. the field is 12 arcmin wide in dec, north is up and east to the left. 0 2000 4000 6000 8000 10000 15.0 14.5 14.0 13.5 13.0 12.5 jd-2,440,000 v2282 sgr asiago-catania figure 2: light curve in the i photographic band of v2282 sgr from the asiago (open squares) and catania (filled squares) archives. due to its southern sky position the coverage shows a marked seasonal bias. 119 roberto nesci et al. 3 ccd observations ccd images of m20 were taken at the cornero observatory1 with a 30cm f/6.5 schmidt-cassegrain telescope and an rc filter in 19 different nights, from june to october 2009. the star showed an oscillating behaviour, with an amplitude of about one magnitude (see fig. 3.) the time interval between a minimum and a maximum may be as short as 10 days, so that the loose time sampling of historic plates may underestimate the actual variability range. furthermore, we obtained with the bfosc instrument at the loiano 1.52 m telescope a b, v and r photometric point on 2011 and an r one in 2012. the star was at b = 17.60, v = 16.31, r = 14.70 on 2011 and r = 14.40 on 2012; both r values are within the variability range of the light curve of 2009. 4 optical spectroscopy with the bfosc instrument we also secured two spectra of the star (on 2011-08-02 and 2012-07-18) at 4 å/pixel dispersion (resolution ∼12 å). the spectra were reduced using the iraf reduction package, following the standard procedures. due to the large airmass no reliable flux calibration could be obtained. the one dimensional spectrum of v2282 sgr shows a number of emission lines which are most probably formed in a region physically close to the star. the two spectra taken at a distance of one year from each other look rather similar. the only remarkable difference is that the [n ii] 6584å line is not apparent in july 2012. figure 3: the light curve of v22822 sgr from ccd images of the cornero observatory. crosses are a comparison star, filled squares are the variable. table 1: adopted comparison stars sequence ra dec id n b v f 270.591 -23.040 a 12.69 13.99 13.37 270.560 -23.068 b 13.52 15.47 14.73 270.522 -23.091 c 12.26 13.15 12.96 12.66 270.503 -23.103 d 13.02 13.72 270.497 -23.096 e 13.49 13.82 270.503 -23.070 f 13.29 14.81 14.44 13.57 270.494 -23.061 g 13.27 13.93 270.486 -23.001 h 12.63 13.15 12.96 12.82 270.633 -23.059 m 12.49 14.40 13.73 270.647 -23.114 n 12.82 13.37 270.637 -23.108 o 12.22 12.44 12.29 12.36 270.612 -23.095 p 12.52 13.71 13.15 12.87 270.692 -22.955 q 13.11 270.696 -22.980 r 13.68 15.36 14.55 270.703 -22.992 s 14.13 270.716 -22.993 t 14.21 14.38 270.685 -23.025 u 14.40 14.52 270.722 -23.067 v 13.89 15.35 14.53 270.689 -23.086 x 13.80 270.685 -23.095 y 12.85 14.42 13.72 13.24 270.696 -23.108 z 13.82 15.40 14.71 table 2: emission lines in v2282 sgr wavelength ion ew (å) 6584 [n ii] -5 6563 hα -20 6548 [n ii] -0.6:: 5007 [o iii] -10 4959 [o iii] -5 4861 hβ -16 4363 [o ii] -6:: 4340 hγ -15 the emission lines in the stellar spectrum are listed in table 2, together with their equivalent width (ew). we remark that doing the sky subtraction from the raw spectrum of the star all the night-sky emission lines disappeared, while the [o iii] and balmer emission lines remained, so we are strongly convinced that they are real. 1http://www.cdso.it 120 v2282 sgr revisited among the absorption lines, the most prominent are the nai d doublet (6.5 å), the mgi b triplet (3.3 å), the blends of several atomic lines centered at 5269 and 6497 å, the cai 4227 å and the g band, indicating an intermediate g-k spectral type. the nai d doublet is very strong, compared to the typical value (2.5 å) of the main sequence or giants g-k stars, suggesting the presence of substantial interstellar/circumstellar matter, as expected in a pms star. figure 4: sed of v2282 sgr from photometric de-reddened (av = 1.3 mag) data. the best-fitting nextgen synthetic spectrum with t=4200 k and log g = 3.5, is over-plotted with a continuous line. to gain information about the stellar photosphere, we fitted the observed sed with low-resolution synthetic spectra calculated from the nextgen model atmospheres [3]. we fixed the star distance and extinction to the values d = 1.7 kpc and av = 1.3 mag reported by [4] and we let teff and log g free to vary, adopting a standard extinction law av = 3.1e(b − v ). this interstellar absorption is consistent with the typical value of av = 0.8 mag/kpc and a distance of 1.7 kpc. the fit was limited to the photometric bands from b to j, because afterwards the photospheric flux could be significantly contaminated by an ir excess. the result of the sed fit is shown in fig. 4 where the optical and near-ir fluxes have been corrected for the extinction and the error bars include the average variation of ± 0.5 mag observed both in r and i bands. the best-fit model is that with log g = 3.5 and teff = 4200 k. with these parameters, a radius of about 4.8 rsun would be derived for the star at the adopted distance. the observed flux appears to be dominated by the stellar photosphere up to the k band, while an ir excess, likely related to a circumstellar disk, is quite evident longward of 3 µm. to quantify how much the derived temperature is sensitive to the adopted av we performed several tests: two extreme cases are reported below. adopting av =0.5 gives teff = 3800 k, log g = 2.5 and a radius 5.1 rsun: such a low temperature is not compatible with the observed spectrum, which shows no molecular bands. adopting a substantially higher value (av =3.0) gives teff = 5600 k, log g = 3.5 and a radius 4.3 rsun: this is still compatible with the observed spectral features but produces somewhat larger residuals than the av = 1.3 case. the derived stellar temperature is clearly an increasing function of the assumed absorption. we remark that the actual absorption is the sum of the interstellar absorption, the circumstellar absorption and the possible contribution from the m20 nebula whose outskirts include the star. a measure of the actual absorption is not easy because the star spectrum is peculiar: in particular the ew of the mgi triplet is low for a g5 star (5600 k), and even more so for a k0 (4400 k) one: it would be typical of a main sequence f5 star, which is too hot for the observed spectral features. this suggests that a strong 121 roberto nesci et al. veiling is present in the spectrum, as indicated also by the presence of strong balmer emission lines. therefore we cannot reliably derive the amount of absorption by comparing the intrinsic b-v with the observed one. the ew of the nai d doublet is 6.5 å, so it is well in the saturated part of the relation between e.w. and e(b-v) by [8] and cannot give an accurate estimate of the color excess, save indicating that it must be substantially larger than the simple interstellar one. 5 infrared observations to better understand the nature of v2282 sgr we considered the j, h and k magnitudes taken at different epochs from 2mass, denis, ukidss catalogues and from [10] (see table 3). from this table it appears evident the near ir variability of the star. table 3: near infrared photometry date j h k note 1998/06/14 12.74 12.02 11.54 2mass 1996/04/01 12.90 ... 11.51 denis 2004/08/23 12.99 12.26 11.93 200inch 2006/07/23 15.71 14.21 12.78 ukidss spitzer/irac images between 3.6 and 8 µm were taken in three different epochs. we performed the photometry of the star using iraf/apphot with the same aperture size of [10]. the star was significantly variable, mainly at shorter wavelengths, enforcing the evidence of its pms nature. a bolometric luminosity, with an uncertainty of about one magnitude, may be computed from the infrared spitzer, jhk, and our optical data. we derived a value of 4.1 lsun in the range from b to k and 24.3 lsun in the range from b up to 24 µm, quite reasonable values for a pms star. 6 conclusions v2282 sgr can be classified as a probable class i premain sequence star associated with an accreting circumstellar disk. its observed optical spectrum is consistent with a late g type star. its long-term irregular optical variability, typical of accreting t tauri stars, can be due to a combination of stellar rotation and circumstellar dust occultation, while the observed transient near-ir excess arises from a combination of variable extinction and changes in the inner accretion disk. its x-ray luminosity (7.6 × 1032erg cm−2 s−1) is compatible with the value for pms with mass ∼ 2 msun [7]. one may wonder if v2282 sgr is actually a foreground object unrelated to the trifid nebula: we think that the strong high-excitation [o iii] and n[ii] lines in its spectrum prove the presence of hot gas in front of the star, which cannot be heated by the low temperature of the stellar photosphere and is therefore related to the m20 nebula. acknowledgement r. nesci thanks the direction of the asiago observatory for hospitality during the scanning of the archive plates. references [1] cambresy, l., rho, j., marshall, d.j., reach, w.t., 2011 a&a 527, 141 [2] glasby j.s., the nebular variables, pergamon press, 1974 [3] hauschildt, p. h., allard, f., ferguson, j., baron, e., & alexander, d. r. 1999, apj, 525, 871 doi:10.1086/307954 [4] lynds, b. t., canzian, b. j., & o’neil, e. j., jr. 1985, apj 288, 164 doi:10.1086/162775 [5] maffei, p., 1963, memsait 34, 441 [8] munari, u., zwitter, t., 1997, a&a 318, 269 [6] ogura, k., & ishida, k. pasj 1975, 27, 119 [7] preibisch t. 2007, memsait 78, 1 [9] rho, j., ramirez, s.v., corcoran, m.f., hamaguchi, k., lefloch, b., 2004, apj 607, 904 doi:10.1086/383081 [10] rho, j., reach, w.t, lefloch, b., fazio, g.g, 2006, apj 643, 965 doi:10.1086/503245 122 http://dx.doi.org/10.1086/307954 http://dx.doi.org/10.1086/162775 http://dx.doi.org/10.1086/383081 http://dx.doi.org/10.1086/503245 introduction photographic light curve ccd observations optical spectroscopy infrared observations conclusions acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0038 acta polytechnica ctu proceedings 4:38–42, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app hydrogen charging of fuel cladding methodology jakub krejčía, ∗, jitka kabátováb, jan kočíb, zuzana weishauptovác, věra vrtílkováb a department of nuclear reactors, faculty of nuclear sciences and physical engineering, czech technical university in prague, czech republic b ujp praha, a.s., prague, czech republic c department of geochemistry, institute of rock structure and mechanics, czech academy of sciences, v.v.i., prague, czech republic ∗ corresponding author: krejci.jakub@seznam.cz abstract. hydrogen content is a very important parameter for mechanical properties of fuel cladding, especially after loca transients. therefore, it is necessary to take into account the amount of hydrogen absorbed in the fuel cladding during normal operation (before a hypothetical loca). the required value of hydrogen content is possible to reach by a long-term pre-oxidation test or a much shorter hydrogen charging experiment. the methodology of hydrogen charging developed in ujp is described in this contribution. results of experiments aiming to prepare samples with uniform hydrides and samples with a rim-layer and other hydrides are shown. keywords: fuel cladding, hydridation. 1. introduction thanks to a low absorption of neutrons and corrosion resistance zirconium alloys are an indispensable material for nuclear fuel cladding tubes. however, their mechanical properties deteriorate during operation in the reactor conditions, where corrosion reactions take place leading to the formation of oxide layers and also to dissolution and precipitation of hydrogen, and the formation of hydrides [1]. zirconium hydrides are brittle, and can initiate crack and fissure formation in the alloy. hydrogen also has a considerable effect on the ductility of the cladding, after a hypothetical loca (loss of coolant accident). much attention has been focused on this topic worldwide [2, 3]. the requirement of sufficient fuel rod strength upon quench taking into account an additional mechanical load is used regarding to the brittle mode, which is associated with cladding oxidation in the presence of steam at high temperature. this requirement led to a new equivalent cladding reacted (ecr) design limit based on loca semi-integral tests developed by jaea (japan atomic energy agency). this criterion is expressed as a function of in-reactor hydrogen pickup [3, 4]. alternative k-criterion and o-beta criterion for e110 alloys were developed in ujp [5, 6] also. both are applicable for specimens with hydrogen content. the presented study shows the methodology of hydrogen charging developed in ujp. samples with uniform hydrides and specimens with rim-layer and other hydrides are shown. double-sided and singlesided hydrogen charging is presented. 2. experimental and methodology 2.1. speciment preparation all samples examined in this study were fabricated of the e110 alloy. the chemical composition is presented in table 1. wt. nb h o hf zr [%/ppm] [%] [ppm] [ppm] [ppm] zr1nb 1.0 3 <400 400 balanced table 1. the chemical composition of studied alloy. the tested tubular specimens had the following dimensions: outside diameter ∼9.1 mm and wall thickness ∼686 µm. for double-sided hydrogen charging 30 mm or 90 mm long non-irradiated segments were used, which were cleaned, degreased, and then weighted, and measured (length, diameter). for single-sided hydrogen charging 90 mm long nonirradiated segments and end-plugs made of a rod from the same material were used. the tube-part and endplugs were welded together using the electron beam welding method. these specimens were cleaned, degreased, and then weighted and measured (length, diameter). after that, the samples were filled with inert gas (argon) at required inner pressure (0–12.5 mpa), and welded in a pressure chamber. the procedure of welding was developed in ujp, and the device can be seen in fig. 1. the samples were weighted after welding, and based on the mass gain of inner gas the inner pressure was recalculated. 38 http://dx.doi.org/10.14311/ap.2016.4.0038 http://ojs.cvut.cz/ojs/index.php/app vol. 4/2016 hydrogen charging of fuel cladding methodology figure 1. device used for-welding samples. 2.2. hydrogen charging using gravimetric sorption analyzer double-sided hydrogen charging was performed in cooperation with the institute of rock structure and mechanics of the czech academy of sciences, and it was published in 2015 [2]. the hydrogen charging was carried out using a microscale of an iga 002 gravimetric sorption analyzer (hiden). a sample was placed into the reactor chamber of the instrument by hanging it vertically on the microscale beam, and the entire system was evacuated at a temperature of 200 ◦c. hydrogen charging was carried out by hydrogen of 99.9999 vol. % purity at a temperature of 405 ◦c and pressure of 0.1 mpa. the mass changes of the sample and the thermodynamic conditions of the process were observed during hydrogen charging. an example of the course of the hydrogen charging process is presented in fig. 2. after the required mass gain was reached, a hydride layer was observed on the surface of the samples with the higher hydride concentration. figure 2. sample packed in aluminum foil. 2.3. autoclave charging an alternative method for hydrogen charging was developed in ujp. the autoclave method provides the possibility to charging bigger amount of samples together. target amount of hydrogen content is up to 600 ppm for pwr conditions (zirlo ∼70 gwd/t [3]), and up to 120 ppm for e110 in vver environment (∼5 years in reactor conditions). specimens were packed in an aluminum foil, which was used as a protection against oxidation, see in fig. 2. the prepared samples were placed in the autoclave with an inner volume of about 4.5 dm3. the hydrogen charging was carried out by hydrogen of 99.9999 vol. % purity at a temperature of 350 ◦c, and pressure of about 0.1 mpa. the temperature and the pressure in the autoclave during the hydrogen charging process are shown in fig. 3. figure 3. temperature and pressure during hydridation process. after the hydrogen charging process as the aluminum foil was removed, a hydride layer and oxide layer was observed on the surface of the samples, see in fig. 4. figure 4. the typical outlook of specimen after hydrogen charging process. samples were weighted after, and based on the mass gain the hydrogen pick-up was calculated. the following form is used for calculation. it presumes, that the whole mass gain corresponds to hydrogen pick-up h = ∆m v0 ∆m v0 + m0 v0 × 106 = 1 1 + m0∆m × 106, (1) where h is hydrogen content in ppm, ∆m is mass gain during hydridation, m0 is initial weight and v0 is sample volume. 39 j. krejčí, j. kabátová, j. kočí et al. acta polytechnica ctu proceedings 2.4. vacuum annealing the dissolving of hydride-layer was processed in a vacuum furnace (clasic). at a pressure about 10−6 bar, and temperatures of about 575 ◦c were used for annealing. the length of the exposure time was chosen up to 6 hours of duration. an influence of the temperature and the exposure time on the hydride-layer dissolution was observed. the surface of the samples after annealing was the same as before the charging process, as can be seen in fig. 5. figure 5. the typical outlook of specimen before and after vacuum annealing process. the samples were weighted after the vacuum annealing, and based on the mass loss the hydrogen pick-up was recalculated. for calculation of hydrogen content is used the equation (1), where ∆m is mass gain after annealing. 2.5. metallographic observations parts of the samples intended for a metallographic evaluation of the hydrides were fixed in epoxy resin, were prepared using a standard polishing procedure, and were etched after. on the prepared metallographic sections, the hydrides were observed and evaluated using a nikon ephiphot 300 light microscope and a lucie image analyzer. 2.6. hydrogen content measurements parts of the samples intended for hydrogen content measurements were analyzed using the analyzer g8 galileo (bruker), which works on the inert gas fusion (igf) principle. 3. results 3.1. uniform distribution of hydrides preparation of specimens with uniform distribution of hydrides was one of the objectives of the presented work. the methodology consisting of autoclave charging and vacuum annealing was applied on the sample ch21. a predicted hydrogen content value – based on weight gain and calculation – was evaluated at 123 wppm. it was in a good agreement with the experimental value gained using the g8 galileo, which was 135 wppm. a satisfactory dissolution of the hydride-layer was observed after 90 minutes of annealing in the vacuum furnace at a temperature of 575 ◦c. figure 6. sample ch21 was prepared using methodology for uniform distribution of hydrides. 3.2. influence of vacuum annealing the gravimetric sorption method was used for doublesided hydrogen of the sample 1h001. this sample was cut into 7 segments which were used for hydrogen content measurements and for experiments focused on the influence of the annealing parameters. the predicted hydrogen content value – based on the weight gain during the hydrogen charging – was evaluated at 540 wppm. it was in a very good agreement with the experimental value gained using the g8 galileo, which was 544 wppm. vacuum annealing conditions for several segments of the sample were described in table 2. segment temperature exposure time figure [◦c] [min] 1h001-6 no annealing no annealing 7 1h001-4 350 20 8 1h001-3 575 20 9 1h001-1 575 60 10 and 11 table 2. the sample 1h001, annealing conditions. the sample 1h001-6 was processed for hydrides evaluation after the hydrogen charging process. hydride layers and a uniform distribution of hydrides in the sample were found (fig. 7). the detail of sample numbered 1h001-4 did not show any observable hydridelayer dissolution (fig. 8). an observable beginning of the hydride-layer dissolution was seen for detail the sample numbered 1h001-3 (fig. 9). the metallographic evaluation of sample numbered 1h001-1 40 vol. 4/2016 hydrogen charging of fuel cladding methodology showed the hydride-layer in an advanced state of dissolution (fig. 10), or the whole hydride-layer dissolved (fig. 11). figure 7. sample 1h001, section 6. prepared after hydrogen charging (without annealing). figure 8. sample 1h001, section 4. after hydrogen charging was used annealing at temperature 350 ◦c and 20 minutes of time duration. figure 9. sample 1h001, section 3. after hydrogen charging was used annealing at temperature 575 ◦c and 20 minutes of time duration. figure 10. sample 1h001, section 1, detail b. after hydrogen charging was used annealing at temperature 575 ◦c and 60 minutes of time duration. figure 11. sample 1h001, section 1, detail a. after hydrogen charging was used annealing at temperature 575 ◦c and 60 minutes of time duration. 3.3. influence of inner pressure the mechanical properties of the fuel cladding depend not only on the hydrogen content, but as well on the orientation of the hydrides. radial hydrides can appear in some specific cases of operation, which can initiate a crack and failure of the fuel cladding. precipitation of radial hydrides was observed during slow cooling of heated samples with inner gas pressure under the temperature of hydrogen dissolution. after hydrogen charging were samples cut into segments. two were used for hydrogen content evaluation and one for metallography. the experimental values of hydrogen content gained using g8 galileo, were 328 wppm (829 447), and 393 wppm (829 463). the number of radial hydrides is directly connected with the value of applied stress, which is caused by the inner gas pressure. figures 12 and 13 show hydrides distributions for several values of inner pressure. an outer hydride-layer was observed in all cases and 41 j. krejčí, j. kabátová, j. kočí et al. acta polytechnica ctu proceedings for samples with inner pressure, especially sample numbered 829 463, radial hydrides were observed. radial hydrides for inner pressure of 0.05 mpa were not observed. the number of radial hydrides was evaluated as 34 % for 829 447 (inner pressure at 20 ◦c – 3 mpa), and 70 % for 829 463 (inner pressure at 20 ◦c – 9 mpa). figure 12. sample 829 447. after hydrogen charging with inner pressure 3 mpa. figure 13. sample 829 463. after hydrogen charging with inner pressure 9 mpa. 4. conclusions the experiments have shown that a requested hydrogen content and a uniform distribution of hydrides can be reached using hydrogen charging, and a further vacuum annealing procedure. influence of temperature and time exposure of vacuum annealing was observed. the experiments have shown the level of hydride-layer dissolution under several conditions. radial hydrides were observed during experiments with inner pressure. a large amount of radial hydrides, about 70 %, were observed for an inner pressure value which can be possibly reached in reactor conditions also. this contribution presented the methodology used for hydrogen charging in ujp. prepared samples is possible to use for loca experiments focused on testing material deteriorated during operation in the reactor conditions. future works will be focused on a more detailed elaboration of several topics. a quantitative evaluation of hydride-layer dissolution and radial hydrides precipitation will be provided. acknowledgements the authors would like to thank to j. šustr, j. sýkora, v. rozkošný, f. manoch, and m. štukbauer for the samples, and metallographic sections preparation, and for the hydrogen content evaluation. financial support for this research through grants no. cz.2.16/3.1.00/21563 and “ta02011025, program alfa – tačr” is gratefully acknowledged. references [1] h.-h. hsu. an evaluation of hydrided zircaloy-4 cladding fracture behavior by x-specimen test. journal of alloys and compounds 426(1–2):256–262, 2006. doi:10.1016/j.jallcom.2005.12.113. [2] z. weishauptová, j. navrátilová, v. vrtílková. the influence of hydrogen on high temperature oxidation of zr1nb cladding tubes. in conference topfuel 2015, proceedings – poster, pp. 319–326. 2015. https://www.euronuclear.org/events/topfuel/ topfuel2015/transactions/ topfuel2015-transactions-poster.pdf. [3] y. yan, t. burtseva, m. billone. recent results at argonne on post-quench ductility of high-burnup cladding. in fuel safety research meeting. japan atomic energy agency, 2009. [4] s. boutin, s. graff. a new loca safety demonstration in france. in conference topfuel 2015, proceedings – part ii, pp. 440–449. 2015. https://www.euronuclear. org/events/topfuel/topfuel2015/transactions/ topfuel2015-transactions-oral-2.pdf. [5] v. vrtílková, l. novotný, r. doucha, j. veselý. an approach to the alternative loca embrittlement criterion. in proceedings of segfsm topical meeting on loca fuel issues. 2004. https://www.oecd-nea. org/nsd/docs/2004/csni-r2004-19.pdf. [6] m. négyesi, v. klouček, j. lorinčík, et al. proposal of new oβ oxidation criterion for new types of the zr1nb alloy of fuel claddings. nuclear engineering and design 261:260–268, 2013. doi:10.1016/j.nucengdes.2012.09.033. 42 http://dx.doi.org/10.1016/j.jallcom.2005.12.113 https://www.euronuclear.org/events/topfuel/topfuel2015/transactions/topfuel2015-transactions-poster.pdf https://www.euronuclear.org/events/topfuel/topfuel2015/transactions/topfuel2015-transactions-poster.pdf https://www.euronuclear.org/events/topfuel/topfuel2015/transactions/topfuel2015-transactions-poster.pdf https://www.euronuclear.org/events/topfuel/topfuel2015/transactions/topfuel2015-transactions-oral-2.pdf https://www.euronuclear.org/events/topfuel/topfuel2015/transactions/topfuel2015-transactions-oral-2.pdf https://www.euronuclear.org/events/topfuel/topfuel2015/transactions/topfuel2015-transactions-oral-2.pdf https://www.oecd-nea.org/nsd/docs/2004/csni-r2004-19.pdf https://www.oecd-nea.org/nsd/docs/2004/csni-r2004-19.pdf http://dx.doi.org/10.1016/j.nucengdes.2012.09.033 acta polytechnica ctu proceedings 4:38–42, 2016 1 introduction 2 experimental and methodology 2.1 speciment preparation 2.2 hydrogen charging using gravimetric sorption analyzer 2.3 autoclave charging 2.4 vacuum annealing 2.5 metallographic observations 2.6 hydrogen content measurements 3 results 3.1 uniform distribution of hydrides 3.2 influence of vacuum annealing 3.3 influence of inner pressure 4 conclusions acknowledgements references 282 acta polytechnica ctu proceedings 2(1): 282–285, 2015 282 doi: 10.14311/app.2015.02.0282 diagnostic of the symbiotic stars environment by thomson, raman and rayleigh scattering processes m. sekeráš1, a. skopal1 1astronomical institute of the slovak academy of sciences, 059 60 tatranská lomnica, slovakia corresponding author: msekeras@ta3.sk abstract symbiotic stars are long-period interacting binaries consisting of a cool giant as the donor star and a white dwarf as the acretor. due to acretion of the material from the giant’s stellar wind, the white dwarf becomes very hot and luminous. the circumstellar material partially ionized by the hot star, represents an ideal medium for processes of scattering. to investigate the symbiotic nebula we modeled the wide wings of the resonance lines ovi λ1032å, λ1038å and heii λ1640å emission line in the spectrum of ag dra, broadened by thomson scattering. on the other hand, raman and rayleigh scattering arise in the neutral part of the circumstellar matter around the giant and provide a powerful tool to probe e.g. the ionization structure of the symbiotic systems and distribution of the neutral hydrogen atoms in the giant’s wind. keywords: symbiotic stars spectroscopy. 1 introduction symbiotic stars are interacting binary systems consisting of a cool giant and a hot compact star most probably a white dwarf. typical orbital periods are between 1 and 3 years, but can be significantly larger. a small part of the material lost by the giant in the form of a stellar wind is accreted by the compact star. this process generates a very hot (th ∼ 105 k) and luminous (lh ∼ 102 − 104 l�) source of radiation, which then ionizes a fraction of the neutral wind from the giant, giving rise to nebular emission in the so-called symbiotic nebula. we distinguish two stages of symbiotic systems. a quiescent phase, when the hot star releases its energy with constant rate and spectral energy distribution, and active phase, when the symbiotic system significantly changes its radiation and brightens up by a few magnitudes in the optical. the partially ionized circumbinary environment represents an ideal medium for scattering processes. in particular we measure effects of the raman and rayleigh scattering of the hot star radiation on neutral atoms of hydrogen and the thomson scattering on free non-relativistic electrons. fig. 1 shows a sketch of simplified ionization structure of symbiotic stars with location of scattering processes in their environment. in this contribution we present the effects of these three types of scattering processes observed in the spectra of symbiotic stars and their possible applications. thomson scattering raman scattering & rayleigh white dwarf red giant ) h ii region (symbiotic nebula h i region figure 1: simplified ionization structure of a symbiotic star. while the radiation from emission region in a vicinity of the white dwarf is thomson scattered in ionized part of the circumstellar environment, in the neutral part it can be raman or rayleigh scattered. 2 thomson scattering the scattering of photons on free low-energy electrons (hν << mec 2) has a very small cross section σt = 6.652 × 10−25cm2 and is wavelength independent. within the symbiotic nebula this scattering will attenuate the cores of emission lines redistributing the scattered part of the line into its wings. the wings be282 http://dx.doi.org/10.14311/app.2015.02.0282 diagnostic of the symbiotic stars environment by thomson, raman and rayleigh scattering processes come very broad, because of a large thermal motions of free electrons, and due to the small value of σt , they will be faint. nevertheless the densest portions of symbiotic nebula with high electron concentration (ne ∼ 108 − 1012cm−3) together with the extremely intense emission lines of highly ionized elements formed in the vicinity of the hot star (e.g. ovi λ1032å, λ1038å) allow us to observe these broad wings. figure 2: a model (solid line) of the observed (dotted line) profile of the ovi λ1032, λ1038å emission lines as a result of thomson scattering in the spectra of ag dra during active and quiescent phase. to model the broad faint wings of the line profile, we adopted the theoretical line profile used already by castor et al. (1970). we applied modeling to the resonance dublet ovi λ1032å, λ1038å and heii λ1640å emission line in the (far-)uv spectra of symbiotic star ag dra. we determined the mean electron optical depth, τe, and electron temperature, te, of the symbiotic nebula during different levels of activity (fig. 2). during quiescent phases the mean value of τe = 0.056 ± 0.006 and te = 19 200 ± 2300 k. on the other hand during active phases τe = 0.64 ± 0.11 and te = 32 300 ± 2000 k (see sekeráš & skopal, 2012 in detail). during the active phases the hot star significantly enhances its stellar wind. this is supported by the broadening of the hα wings (skopal, 2006) as well as by the x-ray/uv flux anticorrelation observed in ag dra during active and quiescent phases (skopal et al., 2009). the material of the wind is ionized by the luminous hot star, which enhances radiation from the symbiotic nebula, which dominates the spectral region u. this produces a significant supplement of free electrons into the nebula. because τe depends on the electron concentration along the line of sight throughout the symbiotic nebula in direction to the hot star, its increase during the active phase indeed indicates the presence of an enhanced mass loss from the white dwarf (fig. 3). 8 9 10 11 0.1 0.3 0.5 0.7 0.9 1.1 1.3 u m ag ni tu de τe figure 3: during the transition from the quiescent phase (u∼11; bottom left) to the active phase (u∼8; top right), the ionized enhanced stellar wind from white dwarf represent a new source of free electrons into the nebula, which increases the electron optical depth τe. 3 raman scattering unlike the thomson scattering on free electrons, which acts in the ionized part of the red giant’s wind (symbiotic nebula), raman scattering arises within its neutral part around the giant. the raman scattering on neutral atoms of hydrogen represents the most pronounced effect. it is a process when a photon excites the atom from its ground state to an intermediate state, which is immediately stabilized by a transition to a different lower main energy level, resulting in emitting a photon of a different frequency. we can observe several raman emission features originated from scattered photons of intense emission lines, such as heii λ6545å (e.g. lee et al., 2003), cii λ7023å, λ7054å (schild & schmid, 1996), heii λ4851å (van groningen, 1993), etc. however the most conspicuous example of raman scattering are broad emission lines at λ6825å and λ7082å (e.g. schmid et al., 1989) (see fig. 4). these are the characteristic features in the spectrum of about half of symbiotic stars. they are result of scattering of resonance dublet of ovi λ1032å, λ1038å in the far-uv on neutral hydrogen atoms. their nature was confirmed by near-simultaneous uv and optical observations and spectropolarimetry (e.g. schild & schmid, 1996, birriel et al., 2000). the raman scattered lines often show a multi-peaked structure, even when the parental ovi emission line is single-peaked. to explain the line profile differences between the ovi lines and their raman scattered counterparts, schmid et al. (1999) suggest doppler shifts introduced in the raman scattering process and non-isotropic ovi line emission. moreover, the evolution of raman lines during quiescent and active phases can reveal the ionization structure around the hot star (e.g. skopal et. al., 2006, fig. 4 here). one of the main characteristic feature of the ramanscattered lines is the broadening according to the 283 m. sekeráš, a. skopal parental uv line. the wavelength λ0 of raman scattered radiation is associated with the wavelength λi of the incident uv radiation, which leads to the broadening of the raman emission lines by factor of λ0/λi. for example the raman emission at 6545å is broadened according to the parental heii λ1025å line by a factor of ≈ 6.4. as raman scattering is a dipol process we can observe strong polarization of the broad emission lines λ6825å and λ7082å. the amount and orientation of the polarization are expected to change periodically due to the orbital motion. this fact can be helpful to provide information concerning the orbital parameters, such as binary period, orientation and inclination of the orbital plane (e.g. schild & schmid, 1996, harries & howarth, 1996). 0 1 2 f lu x [1 012 e rg c m -2 s -1 a -1 ] fuse: 05/07/2002 ϕ = 0.95 optical: 07/08/2002 ϕ = 0.99 o vi λ1032/500 raman λ6825 0 1 2 fuse: 22/10/2002 ϕ = 0.09 optical: 02/02/2003 ϕ = 0.23 o vi λ1032/150 raman λ6825 0 1 2 -200 -100 0 100 200 radial velocity [km/s] fuse: 04/08/2003 ϕ = 0.47 optical: 31/07/2003 ϕ = 0.47 o vi λ1032/400 raman λ6825 figure 4: comparison of the ovi λ1032å line profiles with their raman scattered counterparts along the outburst in symbiotic star z and (adopted from skopal et al., 2006). in addition to polarimetric measurements, the scattering geometry can be estimated by the raman scattering efficiency. it is defined as the ratio of the number of raman scattered photons to the number of emitted far-uv photons. this parameter strongly depends on the size of the neutral scattering region as ’seen’ by the original photons and thus also on the mass loss rate from the cool giant in symbiotic stars (e.g. lee et al., 2003). 4 rayleigh scattering in the case of rayleigh scattering, a photon excites an atom from its ground state to an intermediate state, which is immediately stabilized by a transition to a true bound state and this is followed by immediate reemission of a photon of the same wavelength. thus when the far-uv radiation from/around the hot star goes through the neutral part of the giant’s wind, it can be rayleigh scattered on the neutral hydrogen atoms. effect of this type of scattering is well observed in the form of strong attenuation of the continuum around hydrogen lines of the lyman series. the presence of rayleigh scattering was observed in the spectra of symbiotic stars with a high orbital inclination at/around the position of the inferior conjunction of the giant (e.g. eg and, vogel, 1991). figure 5 demonstrates this case for eg and. 1 3 5 7 9 3.1 3.2 3.3 3.4 3.5 3.6 f lu x [1 013 e rg c m -2 s -1 a -1 ] log wavelength [a] eg and 13/01/1983 ϕ=0.95 figure 5: strong attenuation of the far-uv continuum by rayleigh scattering process (gray area) in the spectrum of eg and, close to the inferior conjunction of the giant. the cross section for raman scattering depends strongly on wavelength with λ−4, which is well demonstrated by monte carlo simulations by schmid (1995), and increases sharply toward lyman resonance transitions. therefore it is easily observable in far-uv spectrum (1200 å 1400 å) and not distinguishable in the optical. in optically thick environment the photon may undergo many rayleigh scatterings until it is raman scattered and escape the neutral region. rayleigh scattering arises in the neutral wind around the giant. if we consider the simplified steady284 diagnostic of the symbiotic stars environment by thomson, raman and rayleigh scattering processes state ionization model of the symbiotic star (see fig. 1), where the hi zone is approximately cone-shaped with the giant at its top (seaquist et al., 1984), it is obvious that the effect of rayleigh scattering will depend on the ionization structure, the inclination of the orbital plane and the orbital phase. it is therefore most easily measurable in eclipsing systems as periodic changes in the far-uv radiation. through the rayleigh scattering we can measure the neutral hydrogen column density in the line of sight as a function of the orbital phase, with typical values of 1020 − 1024cm−2. in this way, it is possible to explore the density and distribution of the neutral hydrogen and thus also the velocity structure in outer atmosphere of a red giant (e.g. vogel, 1991, pereira et al., 1999). 5 conclusions the nature of the environment around the symbiotic stars, with a fully ionized and also the neutral part of the circumstellar matter, allow us to investigate symbiotics through the effects of scattering processes. thomson scattering on free electrons can diagnose the symbiotic nebula by determining the mean electron optical depth, τe, and electron temperature, te. while the nebula is directly connected with the changes of the hot star properties, the effects of thomson scattering can provide some information also about the symbiotic nebula during the quiescent and active phase. raman and rayleigh scattering on neutral hydrogen atoms arises in the neutral part of the giant’s wind. they are observed in a form of broad emission features and an attenuation of the far uv continuum, respectively. these processes are useful to examine the ionization structure of the symbiotic systems, kinematics of the scattering and emitting region, the structure and density distribution of the neutral hydrogen in the giant’s wind and thus its mass loss rate. because the environment, where these scattering processes arises, is a result of the interaction between the giant and the white dwarf, they can provide also information about orbital parameters and physical properties of the two stars independently of other approaches. acknowledgement the research was supported by a vega grant of the slovak academy of sciences no. 2/0002/13. references [1] birriel, j. j., et al.: 2000, apj 545, 1020. doi:10.1086/317851 [2] castor, j.i., smith, l.f., van blerkom, d.: 1970, apj 159, 1119. doi:10.1086/150393 [3] harries, t. j., howarth, i. d.: 1996, a&a 310, 235. [4] lee, hee-won., et al.: 2003, apj 598, 553. doi:10.1086/378886 [5] pereira, c. b., et al.: 1999, a&a 344, 607. [6] schild, h., schmid, h.m.: 1996, a&a 310, 221. [7] schmid, h. m., et al.: 1989, a&a 211, l31. [8] schmid, h. m., et al.: 1999, a&a 348, 950. [9] seaquist, e. r., taylor, a. r., button, s.: 1984, apj 284, 202. doi:10.1086/162399 [10] sekeráš, m., skopal, a..: 2012, mnras 427, 979. doi:10.1111/j.1365-2966.2012.21991.x [11] skopal, a.: 2006, a&a 457, 1003. [12] skopal, a., et al.: 2006, a&a 453, 279. [13] skopal, a., et al.: 2009, a&a 507, 1531. [14] van groningen, e.: 1993, mnras 264, 975. [15] vogel, m.: 1991, a&a 249, 173. 285 http://dx.doi.org/10.1086/317851 http://dx.doi.org/10.1086/150393 http://dx.doi.org/10.1086/378886 http://dx.doi.org/10.1086/162399 http://dx.doi.org/10.1111/j.1365-2966.2012.21991.x introduction thomson scattering raman scattering rayleigh scattering conclusions 238 acta polytechnica ctu proceedings 2(1): 238–241, 2015 238 doi: 10.14311/app.2015.02.0238 spectroscopic monitoring observations of nova v1724 aql in 2012 t. kajikawa1, m. nagashima1, h. kawakita1, a. arai1,2, y. ikeda1,3, m. isogai1,4, m. fujii5, k. ayani6 1kyoto sangyo univearsity / koyama astronomical observatory 2university of hyogo / nishi-harima astronomical observatory 3photocoding 4national astronomical observatory of japan 5fujii-kurosaki observatory 6bisei astronomical observatory corresponding author: kawakthd@cc.kyoto-su.ac.jp abstract spectroscopic and photometric monitoring observations of nova apl 2012 (v1724 apl) were conducted at koyama astronomical observatory, fujii-kurosaki observatory and bisei astronomical observatory. the nova was initially considered as an outbursting pre-main-sequence young stellar object. our monitoring observations have revealed the nova to be a fe ii type classical nova. the temporal evolution of spectra and light curves of the nova were similar to those of a slow nova (e.g., v1280 sco and v5558 sgr). we observed no evidence of molecule formation in v1724 aql in contrast with v2676 oph in which dust formation occurred after the molecular formation in the nova outflow. keywords: cataclysmic variables classical novae optical spectroscopy photometry individual: v1724 aql. 1 introduction nova aql 2012 (later named as v1724 aql) was discovered on 2012 october 20.4 ut (nishiyama & kabashima 2012). the brightness of the nova during the discovery was reported as 12.6 mag. just after the discovery, we conducted our spectroscopic monitoring observations at the koyama astronomical observatory (kao), fujii-kurosaki observatory (fko) and bisei astronomical observatory (bao) in japan. the first optical spectrum taken on october 21.4 ut was reported by fujii (2012) at fko. the object showed a sharp hα emission ( probably with a p cygni profile) on a very red continuum. this object was considered a classical nova affected by a severe interstellar extinction (strong absorption of the na d line was observed). munari (2012) also reported that the spectrum of this object on october 21.8 ut showed a weak and sharp hα emission on a very red continuum, with strong na i and ba ii absorption lines. this spectrum was basically consistent with that reported by fujii (2012). however, it was considered that the spectrum is more similar to that of an outbursting premain-sequence young stellar object than that of a nova before maximum brightness. on 2012 october 23.5 ut, ayani (2012) also reported the optical spectrum of this object taken at bao, with fe ii emission and o i emission with a p cygni profile. this object was finally identified as a fe ii type classical nova according to the classification by williams (1992). here we report the spectroscopic and photometric observations of v1724 aql performed in a collaboration among three observatories (kao, fko and bao) in japan. 2 spectroscopic and photometric observations the observations conducted in our collaboration were follows. spectroscopic observations were performed at three sites: 1. a 0.4-m telescope with a low-dispersion spectrograph (λ = 4500 — 9000 å with r ∼ 500) at fujii-kurosaki observatory, 2. a 1.01-m telescope with a lowand moderatedispersion spectrograph (λ = 3600 — 9000 å with r ∼ 1500 for a low-resolution mode) at bisei astronomical observatory, and 3. a 1.3-m araki telescope with a low-dispersion 238 http://dx.doi.org/10.14311/app.2015.02.0238 spectroscopic monitoring observations of nova v1724 aql in 2012 spectrograph losa/f2 (λ = 3800 — 8000 å with r ∼ 600; shinnaka et al. 2013) at koyama astronomical observatory. in addition to the spectroscopic observations, photometric observations were performed at kao by the imaging camera adler (araki dual-band imager; nakagawa et al. 2013) with intermediate and broadband filters (y-, i′-, and z′-band filters). figure 1 shows the spectral evolution of v1724 aql. figure 2 shows the light curves (with the multi-band light curves taken from aavso database). we also show the close-up views for hα and o i emission lines to make clear the evolution of their line shapes. clearly, the line widths increased for later epochs in the case of hα while the temporal change in the p cygni profile of o i emission indicated faster expansion velocities at later dates. 0 5 10 15 20 25 4000 5000 6000 7000 8000 9000 wavelength [a] +35d:k +26d:k +25d:f +24d:f +22d:f +22d:k +17d:k +16d:f +16d:k +10d:k +8d:k +5d:f +4d:k +4d:b +4d:f +3d:f +2d:f +2d:b +1d:k +0d:f h β f e i i (7 4 ) f e i i (7 4 ) f e i i (7 4 ) o i (7 7 7 4 ) n a i (5 8 8 9 ) m g ii + o i o i( 8 4 4 6 ) f e i i (4 2 ) h α[o i] ( 6 3 0 0 ) c a i i k: kao f: fko b: bao lo g ( f lu x) + c o n st . 1 0 figure 1: low-dispersion spectra of v1724 aql obtained at three sites from 2012 oct 21 to nov 25 ut. all spectra are shifted by different offsets for readability. emission lines were on the highly red continuum. the line width of hα emission became broader and stronger compared with the nearby continuum, and the ca ii triplet emission lines bacame stronger later. note that the bluish continuum might overlap with the spectra of the nova in the shorter wavelength region (λ < 4500 å ), probably due to a forground (or background) star. 239 t. kajikawa et al. 3 discussion and conclusion from the viewpoint of the spectral evolution, the spectra of v1724 aql slowly changed during the period of our observations. emission lines (hα, hβ, fe ii, na i, o i and ca ii) were on the reddened continuum. we estimated e(b − v ) based on the measurements of balmer decrement by the method described in helton et al. (2010), e(b − v )∼3.1, which indicated severe interstellar extinction for the nova. this value is slightly higher than that reported by rudy et al. (2012) but both values indicate severe interstellar reddening. note that the spectra of the nova showed enhancement in a shorter wavelength region (λ < 4500å ) that appeared similar to a bluish continuum component but was noisier. the nova may be overlapped with another foreground (or background) stellar object that had a high temperature. although one may consider that we should estimate e(b − v ) based on the (b − v ) of the nova around the optical brightness maximum, we had to be careful of the contamination with the b-band brightness of the nova. 8 9 10 11 12 13 14 15 16 17 18 19 0 5 10 15 20 25 30 35 40 45 m a g n itu d e s days from 2012 oct 20 ut v r i b y i’ z’ figure 2: optical light curves of v1724 aql obtained at kao with the multi-band light curves of the nova taken from aavso database. the vertical tick-marks at the top of panel indicate the dates of our spectroscopic observations. the line width of hα emission had clearly increased and became stronger compared with the nearby continuum, and o i and ca ii triplet emission lines grew at later dates. the ejection velocity was ∼400 km/s based on the p cygni profile of hα at t=1d after the outburst (figure 3). it then increased; 630 km/s (t=2d), 640 and 750 km/s (t=5d), and 1150 km/s (t=9d) based on the p cygni profiles of o i as shown in figure 4. finally, the ejection velocity reached ∼2000 km/s based on the fwzi of hα (figure 3) after the brightness maximum at t=17d. the ejection velocities were slower during the earlier phase, and faster during later phase. such temporal changes in ejection velocity in v1724 aql was similar to those of slow novae, e.g., v1280 sco (naito et al. 2012) and v5558 sgr (tanaka et al. 2011). v1724 aql was similar to a slow nova from the viewpoint of spectral evolution. 0 0.2 0.4 0.6 0.8 1 -4000 -2000 0 2000 4000 r e la tiv e f lu x velocity [km/s] +10d:k +8d:k +5d:f +4d:k +3d:f +2d:f +1d:k 0d:f 0 0.2 0.4 0.6 0.8 1.0 -4000 -2000 0 2000 4000 r e la tiv e f lu x velocity [km/s] +35d:k +26d:k +25d:f +24d:f +22d:k +17d:k +16d:k +10d:k figure 3: temporal change of hα emission line before and after the brightness maximum at t=17d after the outburst (upper and lower panels, respectively). the line widths became broader at later dates. the black arrow in the lower panel shows the change in blue-shift component that could have been caused by the change in the velocity structure of the ejected materials. based on the light curves shown in figure 2, nova v1724 aql showed very slow evolution in brightness after the outburst until its brightness maximum in the v band (t=17d, v =13.6±0.2 mag). the visual brightness was almost unchanged (but oscillating with small amplitude) until the brightness maximum since the outburst (14.4±0.4 mag in v -band). the evolution in brightness was slow for v1724 aql in earlier phase. however, after the brightness maximum (t=17d), its brightness in v -band become fainter with a decline rate of +0.14 mag/day (t2 was 14.3 days). although this decline rate was indicative of a fast nova rather than a slow nova (payne-gaposchkin 1957), the light curves after the brightness maximum might be affected by dust formation as suggested by the molecular formation observed by rudy et al. (2012). 240 spectroscopic monitoring observations of nova v1724 aql in 2012 -1 -0.5 0 0.5 1 -4000 -3000 -2000 -1000 0 1000 2000 3000 4000 r e la tiv e f lu x velocity [km/s] +10d:k +8d:k +4d:k +1d:k figure 4: temporal change of the p cygni profile for o i (7774 å at rest) emission line before the brightness maximum. the measured ejection velocites based on the p cygni profiles were 630 km/s (at t=2d), 640 and 750 km/s (at t=5d), and 1150 km/s (at t=9d). rudy et al. (2012) reported the detection of co molecular emission in the near-infrared spectra taken on 2012 oct 27 and 28 ut (before the brightness maximum), indicating that molecule formation was ongoing, with dust formation likely to follow. nova v1724 aql was probably a slow nova that showed co emission during the pre-maximum halt (at least, before the maximum). however, there were no hints for c2 and cn absorption bands in optical spectra of the nova (c.f. v2676 oph, nagashima et al. 2014). those molecular absorption bands were expected to appear during several days (as in the cases of v2676 oph and dq her). our spectroscopic observations were performed frequently enough to detect them if a considerable amount of molecules formed in the outflow of the nova. we propose the hypothesis that the ejecta of v1724 aql was oxygen-rich and the atomic carbon might not be over-abundant with respect to the atomic oxygen (c < o) in the ejecta. as demonstrated by the model calculations (pontefract & rawlings 2004), formation of c2 (and also cn) is more difficult than co in an oxygen-rich atmosphere. v1724 aql might have oxygen-rich ejecta. the spectroscopic observations of v1724 aql in the nebular phase (if reported in the future) will be helpful to test this hypothesis. references [1] ayani, k. 2012, electronic telegram no. 3273, central bureau for astronomical telegrams, international astronomical union (ed., green, d.). [2] fujii, m. 2012, electronic telegram no. 3273, central bureau for astronomical telegrams, international astronomical union (ed., green, d.). [3] helton, l.a., woodward, c.e., walter, f.m., vanlandingham, k., schwarz, g.j., evans, a., ness, j.-u., geballe, t.r., gehrz, r.d., greenhouse, m., krautter, j., liller, w., lynch, d.k., rudy, r.j., shore, s.n., starrfield, s., & truran, j. 2010, aj, 140, 1347. [4] munari, u. 2012, electronic telegram no. 3273, central bureau for astronomical telegrams, international astronomical union (ed., green, d.). [5] nagashima, m., arai, a., kajikawa, t., kawakita, h., ktao, e., arasaki, t., taguchi, g., & ikeda, y. 2014, apjl, 780, l26. doi:10.1088/2041-8205/780/2/l26 [6] naito, h., mizoguchi, s., arai, a., tajitsu, a., narusawa, s., yamanaka, m., fujii, m., iijima, t., kinugasa, k., kurita, m., nagayama, t., yamaoka, h., & sadakane, k. 2012, a&a, 543, 86. [7] nakagawa, s., noguchi, r., iino, e., ogura, k., matsumoto, k., arai, a., isogai, m., & uemura, m. 2013, publ. of astron. soc. japan, 65, 7. doi:10.1093/pasj/65.3.70 [8] nishiyama & kabashima 2012, electronic telegram no. 3273, central bureau for astronomical telegrams, international astronomical union (ed., green, d.). [9] payne-gaposchkin, c. 1957, the galactic novae. amsterdam:north-holland. [10] pontefract, m., & rawlings, j.m.c. 2004, mnras , 347, 1294. doi:10.1111/j.1365-2966.2004.07330.x [11] rudy, r.j., laag, e.a., crawford, k.b., russell, r.w., puetter, r.c., & perry, r.b. 2012, electronic telegram no. 3287, central bureau for astronomical telegrams, international astronomical union (ed., green, d.). [12] shinnaka, y., kawakita, h., kobayashi, h., naka, c., arai, a., arasaki, t, kitao, e., taguchi, g., & ikeda, y. 2013, icarus, 222, 734. doi:10.1016/j.icarus.2012.08.001 [13] tanaka, j., nogami, d., fujii, m., ayani, k., kato, t., maehara, h., kiyota, s., & nakajima, k. 2011, publ. of astron. soc. of japan, 63, 911. [14] williams, r. e. 1992, aj, 104, 725. 241 http://dx.doi.org/10.1088/2041-8205/780/2/l26 http://dx.doi.org/10.1093/pasj/65.3.70 http://dx.doi.org/10.1111/j.1365-2966.2004.07330.x http://dx.doi.org/10.1016/j.icarus.2012.08.001 introduction spectroscopic and photometric observations discussion and conclusion acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0029 acta polytechnica ctu proceedings 7:29–32, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app deformation behaviour of gellan gum based artificial bone structures under simulated physiological conditions nela krčmářováa, ∗, jan šleichrta, tomáš fílab, petr koudelkab, daniel kytýřb a czech technical university in prague, faculty of transportation sciences, department of mechanics and materials, konviktská 20, 120 00 prague 1, czech republic b institute of theoretical and applied mechanics as cr, v.v.i., prosecká 76, 190 00 prague 9, czech republic ∗ corresponding author: krcmarova@fd.cvut.cz abstract. the paper deals with investigation of deformation behaviour of gellan gum (gg) based structures prepared for regenerative medicine purposes. investigated material was synthesized as porous spongy-like scaffold reinforced by bioactive glass (bag) nano-particles in different concentrations. deformation behavior was obtained employing custom designed experimental setup. this device equipped with bioreactor chamber allows to test the delivered samples under simulated physiological conditions with controlled flow and temperature. cylindrical samples were subjected to uniaxial quasi-static loading in tension and compression. material properties of plain gg scaffold and reinforced scaffold buffered by 50 wt% and 70 wt% bag were derived from a set of tensile and compression tests. the results are represented in form of stress-strain curves calculated from the acquired force and displacement data. keywords: gellan gum scaffold, reinforcement, uni-axial loading, simulated physiological conditions. 1. introduction the worldwide incidence of bone disorders and conditions have trended steeply upward. especially high income regions are expected twofold increase between 2010 and 2020 [1]. this is the tribute for populations aging coupled with improper nutrient consumption and poor physical activity. globally more than 40 % of women and 30 % of men are at increased risk of emergence of bone disorders [2]. annually in the usa only, more than half a million bone defects are reported. worldwide the treatment cost reaches more than $2.5 billion. the bone disorders treatment using engineered bone tissue has been viewed promising and yet not fully exploited potential alternative to conventional use of autografts and allografts. artificial tissue is overcoming problems with donor site morbidity, loss of bone inductive factors and/or resorption during healing [3]. in general, several essential demands are placed on artificial structure: i) chemical biocompatibility without toxic effect ii) reduction of the stress shielding effect iii) successful diffusion of nutrients and oxygen iv) controlled degradation and resorption [4]. presented paper deals with uni-axial quasi-static testing of artificial spongy-like structure [5] proposed for bone tissue engineering purposes as a bone scaffold. the investigated gellan gum bioactive glass (ggbag) material combines organic (polysaccharitic) component with inorganic (silicon-calcium based) nanoparticles. this approach effectively enables for adaptation of physical and mechanical properties of the synthesized material according to the desired application [6]. the studied material was subjected to quasi-static loading in both tension and compression to evaluate its expected deformation response in interaction with human body. therefore the experiment was carried out under simulated physiological conditions using bioreactor with circulating synthetic plasma. 2. material bioactive-glass-reinforced gellan-gum is a promising material for wide use in bone tissue engineering [7]. originally the gellan-gum (microbial extracted polysaccharide) was used in food and pharmaceutical industry [8]. gg is composed of repeating units consisting of two d-glucose and one of each lrhamnose and d-glucuronic acid [9]. its main advantage is in ability to form highly porous 3d structures when properly cross-linked and fabricated [10]. material investigated in this study was synthesized at jozef stefan institute (slovenia) as porous spongylike structure buffered by bioactive glass (bag) nanoparticles [11]. during the production process gellan gum was dissolved in ultra-pure water by heating the solution for 30 minutes at 90 ◦c. to the hot gg solution a dispersion of bag was admixed and 0.18 wt% cacl2 was added. kept at high temperatures this mixture was then poured into required mould and let there to spontaneously jellify. finally the samples were frozen at −80 ◦c and freeze-dried. 29 http://dx.doi.org/10.14311/app.2017.7.0029 http://ojs.cvut.cz/ojs/index.php/app n. krčmářová, j. šleichrt, t. fíla et al. acta polytechnica ctu proceedings 3. methods for initial awareness about deformation characteristics of synthesized material set of quasi-static experiments was performed. the first goal was to demonstrate possibilities of in house developed experimental infrastructure for this purpose. expected collapse forces was in range of single newtons and precise loading plate positioning was required as well. to obtain more relevant results some modifications of the experimental devices in detail presented in 3.2 was carried out. using this adapted devices basic material properties and stress–strain response were obtained. 3.1. experimental procedure cylindrical samples with height h = 8.6 ± 0.4 mm, diameter d = 5.0 ± 0.1 mm and weight m ≈ 11 mg, ≈ 16 mg and ≈ 24 mg for plain gg scaffold, ggbag reinforced scaffold with 50 wt% and 70 wt% bag respectively were subjected to tensile and compressive loading under wet condition. for wet condition simulating physiological environment of human body infusion solution plasmalyte (bartex, czech republic) was used. loading plate displacement was set typically for 1000 µm corresponding to deformation approx. 11 − 12 % sufficient for significant sample damage. loading rate was set to 2 µms−1. force and position was read-out with sampling frequency 50 sps. 3.2. instrumentation in house developed indentation device for low-force indentation was adapted for tensile and compression test. originally the device was designed using modular aluminum profile (30 × 30 mm) frame bearing i) x and y motorized axis kk40 (hiwin, japan) for sample positioning with repeatability 10 µm ii) indentation axis based on linear stage mgw12 (hiwin, japan) and linear actuator 43 series (haydon kerk, usa) with position accuracy 3 µm and mounting for u9b/c series (hbm, germany) load-cell. this axis was upgraded using the linear actuator with position accuracy 1.5 µm, encoder with resolution 0.5 µm and load cell with nominal force 50 n (the most precise force transducer in u9b/c series). for testing under wet condition testing device was equipped with bioreactor with controlled flow and temperature of circulating fluid. from the heated reservoir is the fluid pumped to the basin surrounding the samples and loading plates. 3.3. strain calculation the investigated material exhibits very low stiffness, which, coupled with high porosity and suboptimal geometry of the samples, induces high potential for significant boundary effects. unfortunately full-field optical strain measurement of the wet samples placed in fluid basin was not possible using the available setup. therefore comparative measurements for contact and contactless strain evaluation methods was performed in previous studies [12, 13]. however production process of the gg-bag samples does not allow to reliably produce cylindrical samples, the diameter of specimens varied in average ±100 µm, and the loaded faces were rough and not plan-parallel strain-stress curve derived directly from force transducer and encoder indicate relevant results. 3.4. stress calculation the stress σ in all experimental analysis was considered as engineering stress obtained using σ = f ac (1) where ac is cross-sectional area of the specimen calculated from minimal sample diameter measured before deformation. force f was acquired by the load-cell. for the purpose of stress calculations, samples were considered ideally cylindrical, neglecting all geometrical irregularities. 4. results material properties and deformation behaviour of plain gg scaffold and gg-bag reinforced scaffold with 50 wt% and 70 wt% bag content were studied in tensile and compression tests under dry and wet conditions. five experiments for each type of material and loading mode were performed. young’s modulus was calculated using linear regression applied on the elastic part of stress–strain diagrams. the calculated elastic properties and yield stresses are listed in tabs. 2–6. sample e [kpa] gg00 1 88.588 ± 0.147 gg00 2 119.507 ± 0.156 gg00 3 126.159 ± 0.142 gg00 4 93.571 ± 0.144 gg00 5 141.089 ± 0.146 gg00 mean 113.783 ± 22.219 table 1. elastic properties of plain gg samples for compression test sample e [kpa] gg00 1 56.343 ± 0.843 gg00 2 33.403 ± 0.209 gg00 3 36.383 ± 0.412 gg00 4 69.077 ± 0.945 gg00 5 70.436 ± 0.951 gg00 mean 53.129 ± 17.562 table 2. elastic properties of plain gg samples for tensile test 30 vol. 7/2017 deformation behaviour of wet gellan gum scaffold sample e [kpa] gg50 1 81.4518 ± 0.145 gg50 2 131.882 ± 0.137 gg50 3 70.8462 ± 0.141 gg50 4 113.892 ± 0.281 gg50 5 140.221 ± 0.201 gg50 mean 107.659 ± 30.528 table 3. elastic properties of gg samples with 50 wt% bag for compression test sample e [kpa] gg50 1 24.406 ± 0.090 gg50 2 27.156 ± 0.085 gg50 3 24.287 ± 0.080 gg50 4 23.719 ± 0.081 gg50 5 21.376 ± 0.089 gg50 mean 24.18 ± 2.061 table 4. elastic properties of gg samples with 50 wt% bag for tensile test sample e [kpa] gg70 1 134.424 ± 0.120 gg70 2 63.566 ± 0.077 gg70 3 79.592 ± 0.078 gg70 4 67.376 ± 0.083 gg70 5 65.131 ± 0.077 gg70 mean 82.018 ± 29.967 table 5. elastic properties of gg samples with 70 wt% bag for compression test sample e [kpa] gg70 1 22.7629 ± 0.080 gg70 2 23.4423 ± 0.082 gg70 3 20.9057 ± 0.081 gg70 4 21.1495 ± 0.101 gg70 5 22.7019 ± 0.079 gg70 mean 22.192 ± 1.105 table 6. elastic properties of gg samples with 70 wt% bag for tensile test all obtained results in form of enveloped stress– strain curves are plotted in figs. 1, 2. the stress– strain area for each type of the scaffold represents minimum and maximum stress for each strain value. 5. conclusion gg-bag samples with 0, 50 and 70 wt% fraction of reinforcing bag particles were subjected to tensile and compressive loading test to evaluate deformation response in simulated physiological condition. it was found out, that the ambient environment has significant influence on mechanical response of the material as the measured properties. the scaffolds wetted by 0 0.002 0.004 0.006 0.008 0.01 0.012 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 c o m p re s s iv e s tr e s s [ m p a ] compressive strain [−] egg00 = 0.113±0.022 mpa egg70 = 0.082±0.029 mpa egg50 = 0.107±0.030 mpa figure 1. stress–strain curves envelope of scaffolds under compressive loading 0 0.005 0.01 0.015 0.02 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 t e n s il e s tr e s s [ m p a ] tensile strain [−] egg00 = 0.053±0.017 mpa egg50 = 0.024±0.002 mpa egg70 = 0.022±0.001 mpa figure 2. stress–strain curves envelope of scaffolds under tensile loading the synthetic plasma solution exhibited radical loss of stiffness, where the elastic modulae decreased more than ten times [12]. no significant reinforcement effect of bag particles was observed during compression test. in case of tensile loading bag buffered samples unexpectedly exhibited lower elastic modulae and ultimate stresses compared to the plain gg samples. that could be given by scaffold cell-wall heterogenity and disintegrity caused by rigid nature of glass particles. it can be concluded that presented analysis proved use of the considered experimental methods together with available experimental infrastructure for the testing of gg based scaffolds. the most limiting part of the experimental setup is load-cell signal-to-noise ratio at desired loading level and generally suboptimal geometrical characteristics of the samples inducing shear stresses during loading. acknowledgements the research was supported by grant agency of the czech technical university in prague (grant no. sgs15/225/ohk2/3t/16), by interreg project com3dxct (atcz38) and by institutional support rvo: 31 n. krčmářová, j. šleichrt, t. fíla et al. acta polytechnica ctu proceedings 68378297. we would like to express our special thanks to ana grantar for sample synthesization. references [1] w. h. organization. world health statistics. who press, 2015. [2] w. h. organization. global recommendations on physical activity for health. who press, switzerland, 2010. [3] a. r. vaccaro, k. chiba, j. g. heller, et al. bone grafting alternatives in spinal surgery. the spine journal 2(3):206 – 215, 2002. doi:10.1016/s1529-9430(02)00180-8. [4] a. r. amini, c. t. laurencin, s. p. nukavarapu. bone tissue engineering: recent advances and challenges. critical reviews in biomedical engineering 40(5):363– 408, 2012. doi:10.1615/critrevbiomedeng.v40.i5.10. [5] l. polo-corrales, m. latorre-esteves, j. e. ramirezvick. scaffold design for bone regeneration. journal of nanoscience and nanotechnology 14(1):15–56, 2014. [6] e. r. morris, k. nishinari, m. rinaudo. gelation of gellan – a review. food hydrocolloids 28(2):373 – 411, 2012. doi:10.1016/j.foodhyd.2012.01.004. [7] m. bououdina. emerging research on bioinspired materials engineering. igi global, 2016. doi:10.4018/978-1-4666-9811-6. [8] d. hoikhman, y. sela. gellan gum based oral controlled release dosage formsa novel platform technology for gastric retention, 2005. wo patent app. pct/il2004/000,654. [9] j. t. oliveira, l. martins, r. picciochi, et al. gellan gum: a new biomaterial for cartilage tissue engineering applications. journal of biomedical materials research part a 93a(3):852–863, 2010. doi:10.1002/jbm.a.32574. [10] n. drnovšek, s. novak, u. dragin, et al. bioactive glass enhances bone ingrowth into the porous titanium coating on orthopaedic implants. international orthopaedics 36(8):1739–1745, 2012. doi:10.1007/s00264-012-1520-y. [11] a. gantar, l. da silva, j. oliveira, et al. nanoparticulate bioactive-glass-reinforced gellan-gum hydrogels for bone-tissue engineering. materials science and engineering c 43:27–36, 2014. cited by 13, doi:10.1016/j.msec.2014.06.045. [12] d. kytýř, t. doktor, o. jiroušek, et al. deformation behaviour of a natural-shaped bone scaffold. materiali in tehnologije 50(3):301–305, 2016. cited by 0, doi:10.17222/mit.2014.190. [13] j. šleichrt, m. adorna, m. neuhäuserová, et al. deformation characteristics of chopped fibre composites subjected to quasi–static tensile loading. acta polytechnica ctu proceedings 3:71–74, 2016. doi:10.14311/app.2016.3.0071. 32 http://dx.doi.org/10.1016/s1529-9430(02)00180-8 http://dx.doi.org/10.1615/critrevbiomedeng.v40.i5.10 http://dx.doi.org/10.1016/j.foodhyd.2012.01.004 http://dx.doi.org/10.4018/978-1-4666-9811-6 http://dx.doi.org/10.1002/jbm.a.32574 http://dx.doi.org/10.1007/s00264-012-1520-y http://dx.doi.org/10.1016/j.msec.2014.06.045 http://dx.doi.org/10.17222/mit.2014.190 http://dx.doi.org/10.14311/app.2016.3.0071 acta polytechnica ctu proceedings 7:29–32, 2017 1 introduction 2 material 3 methods 3.1 experimental procedure 3.2 instrumentation 3.3 strain calculation 3.4 stress calculation 4 results 5 conclusion acknowledgements references 231 acta polytechnica ctu proceedings 1(1): 231–234, 2014 231 doi: 10.14311/app.2014.01.0231 data analysis of globular cluster harris catalogue in view of the king models and their dynamical evolution. i. theoretical model marco merafina1, daniele vitantoni1 1department of physics, university of rome la sapienza, piazzale aldo moro 2, i-00185 rome, italy corresponding author: marco.merafina@roma1.infn.it abstract we discuss the possibility to analyze the problem of gravothermal catastrophe in a new way, by obtaining thermodynamical equations to apply to a selfgravitating system. by using the king distribution function in the framework of statistical mechanics we treat the globular clusters evolution as a sequence of quasi-equilibrium thermodynamical states. keywords: globular clusters gravothermal catastrophe king models thermodynamical stability. 1 introduction globular clusters (gcs) are stellar systems with masses within the interval 104 −106 m�, containing a number of stars of the order of 105. they are considered as nearly spherical systems due to their low values of eccentricity e; at least 50% of gcs have e < 0.1 and there are no clusters with e > 0.2. the core radius rc, namely the radial coordinate at which the brightness becomes one half of the corresponding value at the center of the system, is almost 10 pc, whereas the tidal radius rt, which is the biggest spatial extension of the cluster allowed by the external tidal field, is typically around 50 pc. for their symmetry and age, there is the possibility to test the evolution of a gcs studying a classical single mass king model (king, 1966) in relation to thermodynamical instability phenomena. in fact, in the analysis of the evolution of gcs, stellar encounters strongly contribute in phase space mixing of stellar orbits. in this scenario, thermodynamics plays a centrale role in the gravitational equilibrium and stability of these clusters, being the average binary relaxation time shorter than their old absolute age which ranges between 10 to 13 gyr. this means that fokker-planck approximation, which takes into account the nature of collisions in globular clusters, can determine the distribution function relevant for obtaining the equilibrium configurations of these systems, whereas the tidal effects due to the presence of galactic gravitational potential are responsible of the confination of the cluster. on the other hand, the observations of the luminosity profiles of different gcs (king, 1962) show similar curves depending only on different values of the star concentration, giving the possibility to fit them by an empirical law and suggesting a unique distribution function for the whole sample of clusters (king, 1966). this effect can be described as a change of the main parameters of the cluster during the dynamical evolution (horwitz & katz, 1977), which maintains the form of the distribution like in a sort of reversible trasformation of a gas in thermodynamic equilibrium, also in accordance to numerical simulations existing in literature which result in keeping unchanged the distribution of velocities of stars for a wide range of values of the concentration during the time evolution driven by the fokker-planck equation. therefore, the evolution of globular clusters can be studied by considering small thermodynamic transformations which keep constant the functional form of the velocity distribution of stars like in the framework of boltzmann statistical mechanics. it is important to note that while the equilibrium is given by the form of the distribution which depends on the fokker-planck equation and consider the real nature of collisions, thermodynamics plays a role in the tidal effects acting on the confination of the system, due to a two competitive phenomena: one given by stellar encounters which tend to refresh the tail of high velocities in the distribution and one due to evaporation of stars which prevents the formation of it, maintaining the system in a sort of thermodynamical equilibrium with the same distribution function even if in presence of a cutoff in the velocity of the stars. 2 the effective potential the king df characterizing the energy distribution of stars with the same mass m in a model that describes 231 http://dx.doi.org/10.14311/app.2014.01.0231 marco merafina, daniele vitantoni a spherically symmetric system with isotropic velocity distribution may be written as f(ε) = b [ e−(ε+mϕ)/kθ −e−(ψ+mϕ)/kθ ] for ε ≤ ψ , f(ε) = 0 for ε > ψ . (1) here ψ = m (ϕr −ϕ) is the energy cutoff, corresponding to the maximum kinetic energy that a star can have at a given radial coordinate r, while ϕ is the gravitational potential. this energy is sufficient to reach the border of the equilibrium configuration r = r, being also the difference between the value of the gravitational potential at the edge of configuration and the same quantity evaluated at a generic distance r from the center. the quantity θ is the thermodynamic temperature of the system while b is a constant of normalization. the behavior of equilibrium solutions for king models has been also analyzed by merafina & ruffini (1989) by solving the poisson equilibrium equation in newtonian regime. we can see the presence of a maximum value of the total mass m at increasing values of the central gravitational potential w0, which denotes the arising of a sort of thermodynamic instability at w0 = 1.35 (fig.1). figure 1: mass in function of w0 for families of solutions at different values of the velocity dispersion (merafina & ruffini, 1989). over this value, we can think the system can evolve towards the loss of thermodynamical equilibrium (gravothermal catastrophe), in accordance to the expected evolution of lynden-bell & wood (1968). in order to consider thermodynamic transformations in the framework of statistical mechanics, it is possible to describe the king df like a maxwell-boltzmann one by introducing an effective potential. in this way the evolution of the king models can be treated as a succession of quasi-equilibrium stages by a thermodynamic theory formally equivalent to the classical one. then, the expression of the effective potential is given by φ = −kθ ln [ 1 −e(ε−ψ)/kθ ] (2) and the distribution function can be expressed as f = be−h/kθ , (3) where h = ε + mϕ + φ is the single particle hamiltonian of the system which includes also the gravitational energy of the single star. the effective potential is a screen potential which restricts the phase space of the available velocities for the stars and takes into account the effect of the tidal forces on the system. in this way, the kinetic temperature t connected with the average velocity of the stars, depending on the radial coordinate r, becomes distinguished from the thermodynamic temperature θ, constant all over the equilibrium configuration. from the modified boltzmann df of eq.3, we can deduce the generalized thermodynamical quantities, as the energy u, the thermodynamical pressure π and the entropy s, related to a shell with radial coordinate r. we get n = av ∫ ψ 0 f √ ε dε , (4) u = av ∫ ψ 0 fh √ ε dε , (5) π = 1 3 a ∫ ψ 0 fε3/2 dh dε dε , (6) s = kav ∫ ψ 0 f(1 − ln f) √ ε dε , (7) where we have replaced the costant b with a, being b = aeα/kθ and now f = ae(α−h)/kθ, while α = µ + mϕ is the chemical potential in presence of the gravitational potential ϕ. in this way we can rewrite the first law of thermodynamics and obtain a new form for the eulero expression, that include the extensive and intensive quantities. we can get also an equation of state formally equivalent to classical one which involves the thermodynamical quantities, valid for a shell with radial coordinate r. we have du = θds − πdv + αdn + n〈dh〉 , (8) u = θs − πv + αn , (9) πv = nkθ (10) 232 data analysis of globular cluster harris catalogue... part i and, for kinetic quantities like temperature t and pressure p , pv = nkt . (11) finally, by integrating the expression of u containing the single particle hamiltonian h (eq.5) all over the configuration, we can find an additional term eeff in the expression of the total energy etot of the system, called effective energy etot = ekin + egr + eeff , (12) where ekin and egr are the total kinetic energy and the total gravitational energy, respectively. the partecipation of the effective potential in the total energy corresponds to the account of the tidal potential which determines a finite radius of the cluster. moreover, eq.5 defines the energy of the test shell but, for calculating egr, we need to use the expression egr = 1 2 ∫ r 0 ρϕdv . (13) 2.1 the gravothermal catastrophe thermodynamical instability of a selfgravitating spherical system was first studied by lynden-bell & wood (1968), by considering an isothermal sphere (core) confined in a spherical box. using the classical form of the virial theorem, including a boundary term due to spatial truncation of the density prophile, it is possible, for that system, to calculate the critical value of the central gravitational potential w0 = 6.55 after that thermodynamical instability, known as gravothermal catastrophe, onsets. it is important to note that such instability takes place only in presence of an external thermal bath exchanging heat with the core and driving the system towards the dynamical collapse. with the introduction of the effective potential, we can repeat this analysis for king models and get another critical value for the central gravitational potential w0 = 6.9, which differs from the one obtained by katz (1980) w0 = 7.4, due to the additional term in the total energy etot (see eq.12). the most interesting results concern the profile of specific heat for different values of w0 (see merafina et al., in preparation). by analyzing the behavior of the specific heat all over the configuration, we found different results. the expression of the specific heat cv = (dq/dθ)v arises from eq.8, being constant n and v , by using the expression dq = du −n〈dh〉 . (14) • for w0 < 1.35, we have equilibrium configurations with positive heat capacity all over the system. there are not existing conditions for an evolution of the system towards the critical value corresponding to the onset of the gravothermal catastrophe (w0 = 6.9). further, this particular value (w0 = 1.35) corresponds to one concerning the first maximum mass we found among the equilibrium solutions (see fig.1). • for w0 > 1.35, the system shows an external halo with negative heat capacity and an internal core with a positive value. the system can evolve by increasing the value of w0 until reaching the critical value in which the gravothermal instability onsets. these evolution can take place without the necessity of the presence of an external thermal bath, differently from the previously requested condition in the lynden bell & wood model. results showing the specific heat profiles in function of the radial coordinate for different values of w0 are summarized in fig.2. 0.00 0.25 0.50 0.75 1.00 -0.6 -0.4 -0.2 0.0 0.2 0.4 0.6 0.8 c v/ n k r/r figure 2: behaviour of the specific heat in function of the radial coordinate for different values of w0. 2.2 preliminar observational evidences the stability of the king models was analyzed in detail by katz in 1980, with the same investigation carried out by lynden-bell & wood for the isothermal sphere. katz introduced a new parameter k, which corresponds essentially to the ratio between the escape velocity and the dispersion velocity, both calculated at the center of the cluster. this parameter is directly connected with w0. calculations performed by katz showed that models become thermodynamically unstable over the value k = 8.1 (w0 = 7.4). but, analyzing the sample of data coming from peterson & king (1975) and peterson (1976), katz highlighted an unexplainable gap between the expected value of the sample, k = 7.8 equivalent 233 marco merafina, daniele vitantoni to w0 = 6.9, and the one corresponding to the onset of gravothermal instability, resulting at w0 = 7.4 (see fig.3). figure 3: distribution of galactic gcs at different values of k (katz, 1980). we indeed expect that gcs had enough time to undergo the gravothermal catastrophe and, therefore, the distribution of gcs in terms of k or w0 should peak exactly in correspondence to the critical value. in fact, the primeval gaussian distribution, approaching the critical value during the evolution, deforms in a non-symmetric gaussian curve due to the effect of gravothermal catastrophe which progressively subtracts the collapsed gcs with values of w0 larger than the critical value. for these reasons, the resulting distribution must present a maximum which corresponds to the critical value. it is remarkable to note that, with the introduction of the effective potential in the study of the thermodynamical instability, we obtain a critical value, w0 = 6.9, that bridges this gap and corresponds exactly to the expected value of the sample. this correspondence becomes much more evident by considering the sample of harris (1996) with 127 clusters (if we exclude the pcc ones), as well becomes more evident the non-symmetric form of the distribution. on the other hand, by making a z-test in order to verify the statistical significance of the gap between the stability limit w0 = 7.4 expected by katz and the peak value of the distribution at w0 = 6.9, it can be shown that these two values are not compatible within a confidence level of 95%. 3 conclusions • the additional (positive) contribution of the effective potential on the total energy, considering also the virial condition 2ekin + egr = 0, implies that etot = −ekin + eeff . this enables us to construct models in which the core has a positive heat capacity, allowing to assume the possibility of a survival of the system from the gravothermal cathastrophe which could explain the existence of post core-collapsed objects (pcc). • the model is selfconsistent and admits regions with positive and negative heat capacity which can exchange energy and produce gravothermal instability, without the necessity to assume an external bath as in the lynden-bell & wood model. • we obtain a new critical value for the onset of gravothermal instability by the presence of the effective potential. this value coincides with the value of k (or, equivalently, to w0) corresponding to the peak of the gcs distribution, removing the unexplainable difference outlined by katz. this is an observational evidence of the effects due to the presence of the effective potential, confirmed in the analysis of data of more than 150 gcs contained in the last version of catalogue recently published by harris in 2010 (see also harris, 1996). finally, it may be useful to consider some unsolved problems and perspectives in order to develop the analysis of thermodynamical instabilities of gcs. • the model is not a multimass one and does not take into account the effects in the formation of binary stars. at moment this is a preliminary model which has to be improved. • the new possibility of measuring transverse velocities of the stars in gcs opens important perspectives on the knowledge of the distribution of the star orbits and their eccentricity, in order to better develop n-body simulations in supporting the validity of the model. references [1] cohn, h.: 1980, apj, 242, 765. doi:10.1086/158511 [2] harris, w.e.: 1996, aj, 112, 1487. [3] horwitz, g., katz, j.: 1977, apj, 211, 226. doi:10.1086/154924 [4] katz, j.: 1980, mnras, 190, 497. doi:10.1093/mnras/190.3.497 [5] king, i.: 1962, aj, 67, 471. [6] king, i.: 1966, aj, 71, 64. [7] lynden bell, d., wood, r.: 1968, mnras, 138, 495. doi:10.1093/mnras/138.4.495 [8] merafina, m., ruffini, r.: 1989, a& a 221, 4. [9] merafina, m., fragione, g., piscicchia k.: in preparation. [10] peterson, c.j.: 1976, aj, 81, 617. [11] peterson, c.j., king, i.: 1975, aj, 80, 427. 234 http://dx.doi.org/10.1086/158511 http://dx.doi.org/10.1086/154924 http://dx.doi.org/10.1093/mnras/190.3.497 http://dx.doi.org/10.1093/mnras/138.4.495 introduction the effective potential the gravothermal catastrophe preliminar observational evidences conclusions 146 acta polytechnica ctu proceedings 1(1): 146–150, 2014 146 doi: 10.14311/app.2014.01.0146 origin of x-ray spectral variation and the seemingly broad iron-line spectral feature in seyfert galaxies ken ebisawa1,2, naoki iso1,2, takehiro miyakawa3, hajime inoue1 1institute of space and astronautical science (isas), japan aerospace exploration agency (jaxa), 3-1-1 yoshinodai, chuo-ku, sagamihara, kanagawa 252-5210 2department of astronomy, graduate school of science, the university of tokyo, 7-3-1 hongo, bunkyo-ku, tokyo 113-0033 3tsukuba space center (tksc), japan aerospace exploration agency (jaxa), 2-1-1 sengen, tsukuba-shi, ibaraki 3058505, japan corresponding author: ebisawa@isas.jaxa.jp abstract we present systematic x-ray data analysis of the seyfert galaxies observed by suzaku to study origin of their hard x-ray (2 – 40 kev) variations. in particular, we examine if the “variable partial covering (vpc) model” proposed by miyakawa, ebisawa and inoue (2012), which was successful to explain spectral variations of mcg –6-30-15, is also valid for other seyfert galaxies or not. in this model, intrinsic x-ray luminosity of the agn is not significantly variable, and most observed flux and spectral variations are caused by change of the geometrical covering fraction of the extended x-ray source by ionized absorbing clouds in the line of sight. we found that the observed flux and spectral variations of 20 targets in addition to mcg–6-30-15 are successfully explained by the vpc model. the transmitted spectral component through the absorbing clouds has a characteristics spectral feature of the ionized iron k-edge, which is considered to be the origin of the seemingly broad iron-line feature commonly observed in seyfert galaxies. variation of the partial covering fraction of the constant x-ray luminosity source causes such an anti-correlation between the direct (non-obscured) component and the transmitted (obscured) component, that cancels their variations each other. the cancellation works most effectively at the energy band where intensities of the two components are the closest to each other, namely, just below the iron k-edge. this explains the significantly small fractional variations in the iron k-energy band, another well-known observational characteristic of seyfert galaxies. keywords: black holes active galactic nuclei accretion disk warm absorbers partial covering x-rays individual: mcg–6–30–15. 1 introduction recently, miyakawa, ebisawa and inoue (2012) (mei2012 hereafter) proposed a “variable partial covering (vpc) model ” to explain x-ray spectral variation of seyfert galaxies. in this model, the original xray luminosity of the agn is not significantly variable, and the apparent x-ray variation is primarily caused by variation of the geometrical covering fraction of the extended x-ray source by the intervening clouds having internal ionization structures. mei2012 applied the vpc model to suzaku observations of mcg-6-30-15, and successfully explained not only the small fractional variation in the iron energy band, but also the entire intensity and spectral variations in 1 – 40 kev (see also inoue, miyakawa and ebisawa 2011). it is interesting to examine if the vpc model is also valid for other seyfert galaxies, and intrinsic x-ray luminosities of the seyfert galaxies in general are significantly variable or not. in the present paper, we explore the suzaku archive to select seyfert galaxies which show similar x-ray spectral characteristics to mcg-6-30-15, and apply the vpc model to see if their x-ray intensity and spectral variations are explained by this model or not. 1.1 data selection we chose data from the suzaku public archive, as of 2011 october 5. our main purpose is to study x-ray intensity and spectral variations of the seyfert galaxies, which have similar spectral characteristics to mcg-630-15, in particular in the iron k-energy band. therefore, we selected only targets which are classified as seyfert galaxies, and known to show the seemingly broad iron k-line feature or a hint of that. 146 http://dx.doi.org/10.14311/app.2014.01.0146 origin of x-ray spectral variation and the seemingly broad iron-line spectral feature in seyfert galaxies to study spectral variations efficiently, we chose only the observations which satisfy the following conditions; (1) long enough total exposure time (> 60 ksec) for an observation, (2) bright enough that the total accumulate counts in an observation are more than 50,000 counts in 0.2 – 10 kev, and (3) the sources are variable more than 10% in 4 – 10 kev in an obseravtion. thus, we selected 27 observations for 25 targets (using 50 sequences in total). figure 1 shows two examples of the 0.2–12.0 kev light curves (ngc3516 and ngc4051) among the 27 observations. 2 data analysis and results 2.1 time-averaged spectra firstly, we analyze the time-averaged spectra of the 27 observations from the 25 sources. we apply the “3component model” proposed by mei2012. the model is represented as f = whwl (n1 + w2n2) p + rpn3 + ife, (1) where p is the intrinsic cut-off power-law spectrum, n1 and n2 are the direct power-law normalization and the absorbed power-law normalization, respectively. wh, wl, and w2 represents the transmissions due to highionized warm absorber, low-ionized warm absorber, and partial heavy absorber (nh > 10 24 cm−2), respectively. each warm-absorber has two parameters, the hydrogen column-density, nh, and the ionization parameter, ξ, such that wh = exp(−σ(ξh)nh,h), wl = exp(−σ(ξl)nh,l) and w2 = exp(−σ(ξ2)nh,2), where σ(ξ) means the photo-absorption cross-section. r and n3 are the reflection albedo and reflection normalization by the neutral accretion disk, ife is a narrow iron kα emission line, respectively. the interstellar extinction is also included in the model fitting, but not explicitly shown in equation (1). for details of the spectral model, see mei2012. consequently, we found 22 time-averaged spectra out of 20 sources are successfully represented with the 3-component model. 2.2 intensity-sliced spectra next, we study spectral variations within the 22 observations (out of 20 sources), average spectra of which were successfully fitted by the 3-component model in the previous section. we examine if the ”variable partial covering (vpc) model” proposed by mei2012 may explain the observed spectral variations or not. to prepare, we briefly review the vpc model in the following: let’s define the “total normalization”, n, and the partial covering fraction, α, so that n = n1 + n2, (2) and n1 = (1 − α)n,n2 = αn. (3) detailed study of spectral variations of mcg-6-30-15 by mei2012 suggests that the low-ionized warm absorber, which is optically thin, is associated with the same heavy absorber causing the partial covering (see figure 12 in mei2012), so that wl ≈ 1 − σ(ξl)nh,l ≈ 1 − ασ(ξl)n (fixed) h,l ≈ (1 − α) + α exp(−σ(ξl)n (fixed) h,l ), (4) where n (fixed) h,l is the fixed amount of the column density of the low-ionized absorber associated with the absorbing cloud, presumably as an outer-envelope. consequently, the 3-component model (1) is rewritten as “double partial covering”, such that f = exp (−σ(ξh)nh,h) ×( 1 − α + α exp ( −σ(ξl)n fixed h,l )) × (1 − α + α exp (−σ(ξ2)nh,2)) pn + rpn3 + ife. (5) following mei2012, we create intensity-sliced spectra, and examine the spectral model (5). for each observation, we define four intensity ranges in the 0.2– 12.0 kev light curve so that the photon counts for each intensity range be approximately equal, and create the four intensity-sliced energy spectra. the red dotted horizontal lines in figure 1 indicate the intensity boundaries for each observation. we try to fit the four intensity-spectra in 2 – 40 kev simultaneously only making α variable. as a result, we successfully fitted the intensity-sliced spectra of the 20 observations of 18 sources only varying α. two examples are shown in figure 3. the rest two sources requires n to be slightly variable. 2.3 time-sliced spectra next, we see that time sequences of the energy spectra are also explained by the vpc model with timevarying partial covering fractions. in the xis lightcurves, we define boundaries of the time-sliced spectra every 2.3×104 sec, which corresponds to four suzaku orbital periods. the blue vertical dotted lines in figure 1 show the time-boundaries for each observation, and we made energy spectra from each time bin. numbers of the time-sliced spectra in a single observation are from 6 (4c 74.26) to 28 (ngc 4051). for each observation, a time-series of the spectra are fitted simultaneously with the vpc model in eq. (5). we first try to fit the spectra only allowing the partial covering fraction α to vary. if not successful (i.e., χ2r > 1.2), the total normalization n is varied in addition. 147 ken ebisawa et al. ngc 3516 c o u n t ra te ( s1 ) time (104 s) 0 10 20 30 400 .5 1 .0 1 .5 2 .0 2 .5 ngc 4051 c o u n t ra te ( s1 ) time (105 s) 0 1 2 3 4 0 5 1 0 1 5 2 0 2 5 15 16 figure 1: xis light curves of ngc3516 and ngc4051 in 0.2–12.0 kev. the count rate is binned with 512 s. horizontal red-dotted lines show the count rate intervals with which the intensity-sliced spectra were made (section 2.2). vertical blue-dotted lines show the time intervals with which the time-sliced spectra were made (section 2.3). ngc 3516 in te n si ty ( s1 k e v -1 ) 1 0 -1 1 0 -2 1 0 -3 slice 1 slice 2 slice 3 slice 4 energy (kev) χ 2 5 10 40 -3 0 3 ngc 4051 in te n si ty ( s1 k e v -1 ) 1 1 0 -1 1 0 -2 1 0 -3 slice 1 slice 2 slice 3 slice 4 energy (kev) χ 2 5 10 40 -3 0 3 figure 2: model fitting results of the intensity-sliced xis and pin spectra for ngc3516 and ngc4051. for each observation, the four intensity-sliced spectra are made with approximately equal counts, and the four spectra are fitted simultaneously only varying the partial covering fraction, α. the data and the best-fit power-law model are in the upper panel, while the residuals are in the lower panel. as a result, 21 observations from 19 sources were explained only varying α. figure 3 shows variation of the observed counting rates in 0.2 – 12 kev and the partial covering fraction for two examples among the 21 observations. anti-correlation between the counting rate and the partial covering fraction is obvious, that means the apparent flux increase/decrease of these sources are caused by decrease/increase of the partial covering fraction of the constant luminosity x-ray sources. we notice variability in the optical and uv continuum and lines is often seen to be correlated with observed x-ray variability, while our model requires that the x-ray intrinsic luminosity hardly changes. this suggests that the osculation of the intrinsic x-ray source by the intervening absorbing clouds takes place just inbetween the central black hole and the optical/uv emitting regions, so that the optical/uv lights reflect the variable absorbed x-ray fluxes. 2.4 root mean square(rms) spectra finally, we calculate root mean square (rms) spectra for the 22 observations using the xis data in order to investigate time variation in the iron energy band. we use the same time-series of the spectra used in the previous section as well as the best-fit models. namely, rms spectra are calculated with a bin-width of 2.3 × 104 sec. we divided the time-sliced spectra into 15 energy band, and computed the rms spectra. we show two 148 origin of x-ray spectral variation and the seemingly broad iron-line spectral feature in seyfert galaxies ngc 3516 c o u n t ra te ( s1 ) time (104 s) c o ve rin g fa cto r α 0 10 20 30 40 0 .5 1 .0 1 .5 2 .0 2 .5 0 .0 0 .0 0 .2 0 .4 0 .6 0 .8 1 .0 ngc 4051 c o u n t ra te ( s1 ) time (105 s) c o ve rin g fa cto r α 0 1 2 3 4 0 5 1 0 1 5 15 16 0 .0 0 .2 0 .4 0 .6 0 .8 1 .0 figure 3: variations of the observed xis counting rates (0.2 – 12 kev; black, scale in left) and the partial covering fractions (red, scale in right) for ngc3516 and ngc4051. ngc 3516 v a ri a tio n a m p lit u d e ( % ) energy (kev) 72 5 10 0 1 0 2 0 3 0 4 0 5 0 ngc 4051 v a ri a tio n a m p lit u d e ( % ) energy (kev) 72 5 10 0 1 0 2 0 3 0 4 0 figure 4: rms (root-mean-square) spectra for ngc3516 and ngc4051. black points are calculated from the data, and red-histograms are calculated from the best-fit spectral models for the time-sequence spectra. examples of the 22 rms spectra in figure 4. the small variability of the iron line energy band is successfully explained by the vpc model. references [1] inoue, h., miyakawa, t. & ebisawa, k. 2011, pasj, 63, s669 [2] magdziarz, p., & zdziarski, a. a. 1995, mnras, 273, 837 [3] miller, l., turner, l. j., & reeves, j. n. 2010, mnras, 408, 1928 doi:10.1111/j.1365-2966.2010.17261.x [4] miyakawa, t., ebisawa, k. & inoue, h. 2012, pasj, 64, 140 [5] zoghbi, a. et al. 2010, mnras, 401, 2419 doi:10.1111/j.1365-2966.2009.15816.x discussion matteo guainazzi: (1) it has been recently proposed that the discovery of soft x-ray lags on time-scales consistent with reverberation from a few schwarzschild radii constitute a very strong evidence supporting relativistic disk reflection as a driver of xray variability. can your model reproduce them? (2) can your model put constrains on the location of the warm absorbing clouds? ken ebisawa: (1) we are aware of the recent reports that several seyfert galaxies indicate soft-lags, which some consider due to inner-disk reflection and relativistic reverberation. we have not explored yet how the observed soft-lags are explained in the scheme of our variable partial covering model. [note added in proof: 149 http://dx.doi.org/10.1111/j.1365-2966.2010.17261.x http://dx.doi.org/10.1111/j.1365-2966.2009.15816.x ken ebisawa et al. we noticed that the observed characteristic soft-lag is also consistent with reverberation caused by scattering x-rays passing through highly covering material, as pointed out by miller, turner and reeves (2010). this seems to be a likely mechanism, since our model intrinsically requires rather large covering fractions. however, we also notice that this interpretation does not explain the observed fe kα lags as pointed out by zoghbi et al. (2010).] (2) yes. from spectral variability, we estimate distance from the black hole and sizes of the warm absorbing clouds. please find miyakawa, ebisawa and inoue (2012) and inoue, miyakawa and ebisawa (2011). 150 introduction data selection data analysis and results time-averaged spectra intensity-sliced spectra time-sliced spectra root mean square(rms) spectra acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0102 acta polytechnica ctu proceedings 4:102–106, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app activation analysis of historical samples and neutron spectrum determination at vr-1 training reactor jan šturma∗, milan štefánik department of nuclear reactors, faculty of nuclear sciences and physical engineering, czech technical university in prague, břehová 7, 115 19 praha, czech republic ∗ corresponding author: sturmja2@fjfi.cvut.cz abstract. vertical irradiation channel of the vr-1 training reactor of the department of nuclear reactors ctu was used for activation analysis of historical samples from the 14th to the 19th century. for determination of mass fractions of materials such as copper, gold and silver in irradiated samples, the relative method of activation analysis was used. further, a set of 24 activation detectors of 12 various materials served for the determination of neutron spectrum of the vr-1 reactor using analytical method; moreover the analytical solution was compared to unfolded spectrum obtained from sand-ii deconvolution program. keywords: activation analysis, reaction rate, neutron spectroscopy, historical samples, westcott formalism. 1. introduction research reactors can be used as a significant source of radiation, especially neutrons. there are several ways of neutron applications from medical purposes such as the boron capture therapy to material testing, or transmutation technologies et cetera. one possible application is also a radioanalytical method called the neutron activation analysis. this method is based on intentional production of radionuclides in studied materials through the neutron interactions; it is an exact tool that can be used to qualitatively and quantitatively determine composition of various samples. the pool-type light water training reactor vr-1 of the department of nuclear reactors ctu is frequently used for purposes of neutron activation analysis especially due to its experimental equipment such as vertical and horizontal irradiation channels. presented project aims to determine the composition of historical samples and neutron spectrum of vr-1 training reactor by using the neutron activation technique. 1.1. neutron activaton analysis measuring the characteristic radiation of the sample, usually gamma or beta particles, produced in radionuclide decays, the saturated activity of irradiated sample can be determined as asat = λtreal s(eγ) tlive εfep(eγ) iγ(eγ)( 1 − e(−λta) ) e(−λtc) ( 1 − e(−λtreal) ) ∏ i ci, (1) where λ is the decay constant of measured radionuclide, treal is the real time of the gamma spectrum measurement, tlive is the live time of the measurement, s(eγ) is number of counts in the full energy peak at energy eγ, εfep is the detector efficiency for full energy peak, iγ is the probability of gamma emission, ta is the irradiation time and tc is the cooling time. possible corrections, e.g. the correction for reactor power fluctuation, self-shielding of gamma radiation or the correction for gamma coincidences, are represented by factors ci. for determination of sample composition, the relative method of the activation analysis is standardly used. studied samples are irradiated together with the activation standards of known composition. the amount of investigated material in the sample can be determined as follows ast mst = asp msp , (2) where ast and asp are the saturated activities of produced radionuclide in standard or in sample, and mst and msp are the masses of investigated material in standard foil or in sample. other technique of neutron activation analysis is the absolute method, where the amount of material in sample is defined as msp = aspm naφσa , where m is the molar mass of examined material, na is the avogadro’s number, φ is the integral neutron flux and σa is the spectrum averaged microscopic activation cross-section. 1.2. neutron spectrum determination by activation detectors methods of neutron activation analysis can be also used for measurement of neutron flux using activation detectors of various materials. the response of the 102 http://dx.doi.org/10.14311/ap.2016.4.0102 http://ojs.cvut.cz/ojs/index.php/app vol. 4/2016 activation analysis and neutron spectrum determination at vr-1 reactor detector to the neutron field per one nuclei, called reaction rate, can be calculated as rr = asat n0 , (3) where n0 is the number of target nuclei in the activation detector. reaction rates obtained by irradiating the set of various detectors together with the cadmium cover as well as some threshold interactions make possible to reconstruct the neutron field using analytical methods. the analytical expression of neutron flux is φ(en) = φt(en) + φe(en)∆1∆2 + φf(en). (4) neutron flux in thermal energy region, φt(en), is for well moderated reactors defined according to the westcott formalism as [1] φt(en) = en (ktn)2 e− en ktn rtr g(tn)σ0 , where en is the energy of neutrons, k is the boltzmann constant, tn is the thermodynamic temperature of neutrons, rtr is the response of detector to thermal neutrons, g(tn) is the deviation of the activation cross section from the law 1/v in thermal energy region and σ0 is the activation cross-section for energy en = 0.0253 ev. the activation detector in cadmium cover is activated mostly by neutrons of energies associated with the energy of most significant resonance in crosssection, and therefore it is suitable to determine the epithermal neutron flux. discrete values of neutron flux for resonance energies eres of activation materials are defined as [2] φt(eres) = r epi r k∫e2 e1 σ(en) den , where repir is the response of the detector in cadmium cover to the neutron field, k is the proportional representation of the most significant resonance in the cross-section in epithermal energy region, and e1 or e2 represents the lower or upper boundary of resonance. experimentally obtained discrete values of neutron flux can be fitted using the expected shape of epithermal reactor spectrum in the form of a/exn, and consequently it can be smoothly connected to the thermal and fast analytical solution of neutron spectra using the joining functions ∆1 and ∆2 defined as ∆1 = 1 1 + ( 4.95ktn en )7 , ∆2 = exp( − en3 · 105 ) . threshold activation detectors are serving to express the reaction rates rfr of activation detectors to the fast neutron region. in fast region, the differential neutron flux is described as φf(en) = rfr σt s(en), where σt is the spectrum averaged threshold crosssection and the s(en) is the fission neutron spectrum. 2. materials and methods 2.1. activation experiments of historical samples the vertical irradiation channel of vr-1 training reactor was used to perform the activation experiments of historical coins; the main purpose was to determine the qualitative and quantitative composition by employing the relative method of neutron activation analysis. in 2014, six historical coins were irradiated together with the reference standards of expected materials. in particular, kurush-1923, pfennig-1745, kreutzer-1843, parvus-14.century, centessimo-1852, and heller-1854 were studied, and copper, gold, silver, and nickel foils were used as activation standards in two sets. in 2015, another three coins (korona-1893, kreutzer-1861, and pfennig-1776) together with the silver and copper standards in the form of foils were irradiated. figure 1. first set of historical samples with reference materials. figure 2. third set of historical coins analyzed using activation analysis. figure 3. hpge detector in pb shielding. each set of samples was irradiated twice. the first irradiation at low reactor power was performed to disprove the presence of inappropriate isotopes, such as 59co. because after irradiation of such isotope, the long term radionuclide is produced, and it is inconvenient. gamma lines of 116min were observed in the spectrum of one sample, and due to this fact, the 103 jan šturma, milan štefánik acta polytechnica ctu proceedings indium reference activation standard was added to the set before main irradiation at the nominal power. during the second irradiation, the reactor was operated at power-level of 1 kw, and the irradiating time was set to 15–30 min. after the main irradiation, the gamma-ray spectra of irradiated samples were measured by the semiconductor hpge detector. for all radionuclides observed in spectra, the saturated activity was determined by using the equation (1). 2.2. irradiationg of activaction detectors altogether 24 activation detectors made of 12 various materials (al, ag, au, cu, fe, in, lual, mn, mo, nacl, u and w) in the form of thin foils (nacl detector had a form of tablet) were irradiated in the vertical irradiation channel at mid-height of reactor core of vr-1 reactor. irradiating time was 15–45 min at the reactor-power of 1 kw. activation materials were chosen with respect to the activation cross-sections in order to cover all energy regions of reactor spectrum; therefore the cadmium cover and some threshold detectors were used. after irradiation, the gamma spectra of detectors were measured. for each observed reaction, the reaction rates were determined using the expression (3). figure 4. set of activation detectors. reaction t1/2 rr [s −1] dev [s−1] 197au(n,γ)198au 2.70 days 1.77 · 10−13 4.04 · 10−15 197au(n,γ)198au cd 2.70 days 3.70 · 10−14 2.61 · 10−16 63cu(n,γ)64cu 12.70 hours 7.89 · 10−15 1.62 · 10−16 63cu(n,γ)64cu cd 12.70 hours 3.51 · 10−16 5.50 · 10−18 186w(n,γ)187w 23.72 hours 9.05 · 10−14 7.77 · 10−16 186w(n,γ)187w cd 23.72 hours 1.56 · 10−14 1.20 · 10−16 98mo(n,γ)99mo 65.94 hours 8.46 · 10−16 2.12 · 10−17 98mo(n,γ)99mo cd 65.94 hours 5.61 · 10−16 1.46 · 10−17 115in(n,γ)116min cd 54.29 min 4.06 · 10−14 2.13 · 10−16 115in(n,n’)115min 4.47 hours 1.44 · 10−16 3.76 · 10−18 115in(n,γ)116min 54.29 min 2.19 · 10−13 1.79 · 10−15 238u(n,γ)239u 23.45 min 8.70 · 10−15 3.23 · 10−17 238u(n,γ)239u cd 23.45 min 4.99 · 10−15 1.45 · 10−17 23na(n,γ)24na cd 14.96 hours 2.64 · 10−17 4.01 · 10−19 27al(n,p)27mg 9.46 min 7.77 · 10−18 3.12 · 10−19 27al(n,α)24na 14.96 hours 5.29 · 10−19 4.12 · 10−20 56fe(n,p)56mn 2.58 hours 1.10 · 10−18 2.72 · 10−20 109ag(n,γ)110mag cd 249.79 days 1.79 · 10−15 7.59 · 10−18 55mn(n,γ)56mn cd 2.58 hours 9.25 · 10−16 6.14 · 10−18 55mn(n,γ)56mn 2.58 hours 2.37 · 10−14 5.28 · 10−16 176lu(n,γ)177lu cd 6.73 days 6.89 · 10−14 1.85 · 10−15 176lu(n,γ)177lu 6.73 days 5.70 · 10−12 5.05 · 10−14 23na(n,γ)24na 14.96 hours 7.65 · 10−16 7.82 · 10−18 37cl(n,γ)38cl 37.24 min 7.01 · 10−16 9.07 · 10−18 table 1. reaction rates of irradiated activation detectors. 3. results 3.1. analysis of historical coins composition the relative method of neutron activation analysis was used to determine the quantitative composition of five materials in all irradiated historical coins using equation (2). measured mass and mass fractions wi are listed in table 2. sample material m [mg] wi [%] kurush au 1 052.64 ± 12.54 58.44 ± 0.70 pfennig ag 17.97 ± 0.16 5.62 ± 0.05 (1745) cu 265.72 ± 5.49 83.17 ± 1.72 kreutzer cu 3 682.40 ± 96.71 94.95 ± 2.49 (1843) parvus ag 49.11 ± 0.43 13.67 ± 0.12 cu 194.81 ± 4.65 54.22 ± 1.29 centessimo cu 1 312.61 ± 11.46 99.95 ± 0.87 in (12.34 ± 0.58)·10−3 (76.10 ± 4.42)·10−5 heller ag 2.01 ± 0.31 0.04 ± 0.01 cu 4 797.12 ± 77.15 93.83 ± 1.51 pfennig ag 37.48 ± 0.99 12.82 ± 0.34 (1776) cu 265.88 ± 6.71 90.96 ± 2.29 korona ag 2 674.09 ± 3.097 54.98 ± 0.64 cu 448.64 ± 11.9 9.22 ± 0.42 kreutzer cu 2 889.20 ± 71.95 92.89 ± 2.31 (1861) table 2. determined composition of irradiated coins [3]. parameters of radioactive decays from [4], such as the half-life period, energy and intensity of gamma radiation, were used to assign the observed gamma lines in measured spectra to specific radionuclides. then according to the neutron energies reached in thermal nuclear reactors, cross-sections, and natural isotopic composition, the observed radionuclides were assigned to considered materials. qualitative analysis of composition was carried out in this way for the coin heller. except the contribution of natural background, the gamma spectrum of heller (see fig. 5) contains the gamma lines of 64cu, 110mag marked in blue color, and then 122sb, 124sb, 76as and 24na as well. observed radionuclides referred to presence of antimony, arsenic, mercury, silver and sodium in irradiated samples. figure 5. analyzed gamma spectrum of heller. 104 vol. 4/2016 activation analysis and neutron spectrum determination at vr-1 reactor 3.2. neutron spectrum reconstruction reaction rates of interaction observed in used activation detectors through produced radionuclides covered all the energy regions of neutron spectra for typical thermal reactor. the use of cadmium cover enabled to separate the contributions of thermal and epithermal neutrons. threshold interactions served to determine the response of detector to the fast neutron. using nuclear data from endf/b-vii.1 [5], the neutron spectrum was determined according to equation (4). the irradiation position in the fuel cell dummy with the overmoderated neutron spectrum allowed to utilize the westcott formalism. separately described energy regions of neutron spectrum together with the discrete points of neutron flux, associated with the most significant resonances in activation cross-sections, and total neutron flux, determined using analytical methods, are depicted in fig. 6 and fig. 7 respecitvely. figure 6. analytic description of neutron flux regions. figure 7. total neutron flux determined by analytical methods. results obtained by analytical methods were compared to numerical solution using the unfolding procedure of sand-ii code [6]. as the input (guess) neutron spectrum requested for the deconvolution process, the analytical spectrum was used. in the 5th iteration and using nuclear data from the endf/bvii.1 data library and experimentally determined reaction rates, the neutron spectrum in 36 energy bins from 10−12 mev to 101 mev was successfully unfolded. comparison of determined numerical and analytical neutron spectrum in the mid-height of reactor core of vr-1 training reactor is in fig. 8. calculated ratios c/e, representing calculated over experimental reaction rates rations, as a parameter of spectrum unfolding, are listed in table 3. figure 8. comparison of analytical and unfolded spectrum. reaction c/e [–] reaction c/e [–] 197au(n,γ)198au 0.92 56fe(n,p)56mn 0.99 63cu(n,γ)64cu 0.87 55mn(n,γ)56mn 0.84 115in(n,γ)116min 1.18 176lu(n,γ)177lu 1.06 23na(n,γ)24na 1.03 37cl(n,γ)38cl 0.92 27al(n,α)24na 1.06 238u(n,γ)239u 1.01 table 3. c/e ratios of reactions used for the spectrum unfolding. 4. conclusions according to the experiences and results of activation experiments, the vr-1 reactor is suitable for analyzing of composition of historical samples, especially coins. table 2 clearly shows the possibilities of determination of all materials commonly used in coin production, such as gold, silver, or copper. main advantages of vr-1 training reactor is the possibility of easy way of inserting samples into irradiating position in vertical channel and relatively low reactor power, comparing to others research reactors in czech republic, so it is hard to produce emitters of very high activities. the first time used method of analytical spectrum description in full energy range showed the new possibility of neutron spectrum determination using the activation detectors at vr-1 reactor. moreover, the analytical spectrum corresponds to expectations, but compared to results of numerical solution using the sand-ii unfolding code appears slightly undervalued in the thermal energy region. this brings a new way of potential further specifying of the analytical spectrum description. results illustrated in this paper are a summary of outcome achieved within the bachelor’s thesis in 2013– 2014 and the master’s degree project in 2015–2016. references [1] c. westcott, w. walker, t. alexander. effective cross sections and cadmium ratios for the neutron spectra of thermal reactors. tech. rep., atomic energy of canada ltd., 1959. conference: 2. united nations international conference on the peaceful uses of atomic energy, geneva (switzerland), 1958. http://www.osti.gov/scitech/biblio/4279663. [2] h.-l. pai, p.-y. ma, s.-c. wang, w. s. lee. measurement of epithermal neutron spectra by 105 http://www.osti.gov/scitech/biblio/4279663 jan šturma, milan štefánik acta polytechnica ctu proceedings resonance detectors. nuclear science and engineering 9(4):519–520, 1961. [3] j. šturma. neutronová aktivační analýza historických předmětů na školním reaktoru vr-1. bachelor’s thesis, kjr fjfi, české vysoké učení technické v praze, 2014. [4] s. chu, l. ekström, r. firestone. the lund/lbnl nuclear data search, version 2.0, february 1999. http://nucleardata.nuclear.lu.se/toi/. [5] national nucelar data center. endf/b-vii.1 evaluated nucelar data library. http://www.nndc.bnl.gov/endf/b7.1/. [6] p. griffin, j. kelly, j. vandenburg. user’s manual for snl-sand-ii code. tech. rep. sand93-3957, sandia national laboratories, 1994. http://prod.sandia. gov/techlib/access-control.cgi/1993/933957.pdf. 106 http://nucleardata.nuclear.lu.se/toi/ http://www.nndc.bnl.gov/endf/b7.1/ http://prod.sandia.gov/techlib/access-control.cgi/1993/933957.pdf http://prod.sandia.gov/techlib/access-control.cgi/1993/933957.pdf acta polytechnica ctu proceedings 4:102–106, 2016 1 introduction 1.1 neutron activaton analysis 1.2 neutron spectrum determination by activation detectors 2 materials and methods 2.1 activation experiments of historical samples 2.2 irradiationg of activaction detectors 3 results 3.1 analysis of historical coins composition 3.2 neutron spectrum reconstruction 4 conclusions references 81 acta polytechnica ctu proceedings 2(1): 81–85, 2015 81 doi: 10.14311/app.2015.02.0081 activity of the polar am her (rx j1816.2+4952): a short review v. šimon1,2, a. henden3 1astronomical institute, academy of sciences of the czech republic, 25165 ondřejov, czech republic 2czech technical university in prague, fel, prague, czech republic 3aavso, 49 bay state road, cambridge, ma 02138, usa corresponding author: simon@asu.cas.cz abstract we show that am her displays the transitions between the high and low states with an intermittently existing dominant cycle with length between 400 and 800 days. moreover, these transitions accumulate in clusters, which produces an additional long cycle after smoothing; a single isolated short episode of the low state does not suggest a break of this cycle. the seasons of existence of the cycle can be controlled by the lifetime of the active regions (e.g. prominences, spots) on the donor. in some high-state episodes, a higher luminosity of the bremsstrahlung emission is not accompanied by a higher optical (cyclotron+stream) emission. part of the bremsstrahlung emission can be buried in some episodes. changes of the structure of the accretion region(s) are necessary to explain the variations of the optical and x-ray activity in the high-state episodes of am her. keywords: cataclysmic variables polars activity emission mechanisms optical x-rays individual: am her. 1 introduction am her is the prototype of polars (e.g. [19]). its activity is dominated by the large-amplitude long-term variations: alternating high and low states of the optical brightness usually on the timescales of months and years (e.g. [8, 21, 10]). a small accreting region (about 0.1 of the radius of the white dwarf (wd)) near the magnetic pole of this object is the source of radiation and polarization of polar via several processes: cyclotron (accretion column – mainly optical and ir emission), thermal (soft x-rays – surface of the wd heated by the accreting region), bremsstrahlung (accretion column – hard x-rays) (e.g. [19, 15]). the conditions in this region and the mass accretion rate play an important role in governing the observed activity. 2 long-term activity in the optical band fig. 1a shows the long-term activity of am her in the optical band spanning about 32.7 years. the data come from the aavso international database (massachusetts, usa) [6]. the highly variable orbital modulation has a large amplitude mainly in the high state (e.g. [9, 10]), so the character of the long-term activity is less discernible. fig. 1a therefore shows the hec13 fit running through the orbital modulation (each point is a smoothed value of brightness in the night of observing).the division between the high and low states is 14.3 mag(v ). the properties of the statistical distribution of brightness of the high states strongly evolve with time (fig. 1cd). the typical high-state brightness is near the middle of the distribution on the magnitude scale in most time segments. the low states often show skewness of brightness toward the brighter magnitudes. the fuzzy bright edge of the high-state brightness, smoothed over the orbital modulation, in the histogram suggests that the high state is not a uniquely defined level of luminosity of am her (fig. 1cd). it is questionable whether any “unperturbed” state of the system exists et all. a typical strength of the active regions (large loops of magnetically confined gas near the donor due to its magnetic activity [9] and star spots [13, 7]) during the high state is therefore needed (the times of the quite quiescent donor are thus very rare). in the scenario proposed by [21], this statistical distribution also suggests a typical configuration of the magnetic field of the whole system (donor+wd), around which the system balances during the high state of activity. the brightness of the low states in am her often stabilizes at the baseline level. the mass transfer via roche lobe overflow almost ceases, only clumpy accretion onto the wd from the donor’s wind remains [3, 9]. the strong skewness of brightness in fig. 1cd can emerge if the hec13 fit runs through the variably recurring brightenings. 81 http://dx.doi.org/10.14311/app.2015.02.0081 v. šimon, a. henden moving averages (ma) of brightness with various filter half-widths q enable to search for the evolution of the optical activity of am her on timescales of several years, significantly longer than the duration of the individual episodes of the high or low states (fig. 1). the transitions between these states tend to occur in clusters in am her. smoothing then reveals a long cycle because of the variable clustering of the episodes of the low states. since the borders of these clusters are fuzzy, a smooth profile of the cycle results when ma are applied. a single isolated short episode of the high or the low state does not suggest a break of this cycle since the cycle is caused by the repeating accumulations of these transitions in clusters. some long cycles thus may not be easily discernible without investigating the variations of the parameters of activity, which ma enable (fig. 1ab). figure 1: (a) the long-term activity of am her in the optical band (aavso data [6]). hec13 fit runs through the orbital modulation. the mapping is divided into segments. the smooth lines represent the moving averages (ma). (b) residuals of the ma. (c, d) examples of the strikingly different statistical distributions of brightness. see sect. 2 for details. search for the cycle-length in the transitions between the high and low states from fig. 1a was carried out also with the weighted wavelet z-transform (wwz) developed by [1]. this method enables one to determine the period and amplitude of unevenly sampled time series. wwz indicates whether or not there is a periodic fluctuation at a given time at a given frequency f (fig. 2a). the wwz-transform really finds that the transitions between the states are not quite chaotic. a dominant cycle in the transitions really exists. however, the cycle-length of the transitions is not stable and this cycle is not conclusively detected in some time segments. fig. 2b shows the best cycle-length, determined from f that has the biggest value of wwz at a given time (only the segments in which the amplitude is larger than 60 percent of its peak value in the investigated range of f). in summary, this cycle is characteristic only for a limited time segment. it is intermittent and unstable in length. moreover, the wwz-method shows also another temporarily existing shorter cycles in fig. 2a. for comparison, the cycles of 178 d, 1836 d, 830 d and 1520 d were found by fourier analysis by [11]. the differential rotation of the lobe-filling donor (modeled by [18]) can influence the position of the above-mentioned active regions on the donor with respect to the l1 point. the breaks of the cycle of the transitions between the high and the low states in am her can be caused by a limited lifetime of the above-mentioned active regions connected with the activity of the donor. figure 2: (a) wwz-transform of the light curve of am her from fig. 1a. (b) the best cycle-length. see sect. 2 for details. 3 optical and hard x-ray intensities in the high states in am her, increase of the mass transfer rate from the donor (giving rise to a high-state episode) also establishes specific division of the emission into various spectral regions during the accretion process (fig. 3) (see our analysis in [17] for details). the properties of the emitting region(s) on the wd are established in the 82 activity of the polar am her (rx j1816.2+4952): a short review early phase of the high-state episode (but not reproduced for every episode) of am her. a higher luminosity of the bremsstrahlung emission may not be always accompanied by a higher optical (cyclotron+stream) emission in a given episode of the high state (the intensities in both of these bands were averaged over the orbital period) [17]. this can significantly differ even for two consecutive episodes (fig. 3bc). this relation of intensities is representative for the whole such episode. purely geometric effects (asynchronous rotation of the wd, precession of the wd spin axis) cannot explain the observations – gradual evolution of the x-ray and optical intensities with time would be needed (see [17] for details). we propose the following scenario. the source of the bremsstrahlung emission is confined to a smaller region than that of the cyclotron emission. part of the bremsstrahlung emission is buried in some high-state episodes. since there is no smooth transition of the ratio of the optical and hard x-ray intensities [17], it suggests that the role of several modes of accretion is worth considering. figure 3: (a) part of the asm/rxte (1.5–12 kev) [12] light curve of am her (15-d means). the mean levels and their standard errors are marked. (b, c) intensities of the optical (afoev data, the intensity i = 1 for 14.3 mag) and hard x-ray (1.5–12 kev) (asm/rxte data) emissions in two consecutive episodes of the high state. a higher luminosity of the hard x-ray emission may not be always accompanied by a higher optical emission in a given episode. 4 role of the accretion modes the orbital modulation of am her displays significant changes in both the optical and x-ray spectral regions. unexpected mode of the x-ray (soft: 0.1–0.31 kev, hard: 1.9–8.5 kev) orbital modulation was observed in two observing runs of the same optical high state by [4]. it suggested a two-pole accretion mode in which the primary pole was the dominant emitter of the hard x-rays while the secondary pole was dominant in the soft x-rays. the accretion mode is representative only for a given high-state episode of am her [14]. a short episode of the high state (but less bright than most other highstate episodes) displayed a single-pole accretion mode with a self-occultation of the primary pole, resulting in a deep primary minimum in the bands between e = 0.6 and 10 kev, but no sign of the secondary minimum. the subsequent brighter high state displayed a two-pole accretion mode with a shallow primary minimum in the 2– 30 kev band, suggesting the big dimensions of the accretion region. the secondary minimum occurred only in the soft band (1–2 kev), suggesting that the secondary pole is the dominant emitter of the soft x-rays. the same situation as those observed in one of the previous high states by [4], that is the primary pole dominant in the hard x-rays and the secondary one dominant in soft x-rays, thus repeated. also cyclotron emission undergoes peculiar changes with the accretion mode. in the above-mentioned bright high state observed by [14], cyclotron emission from a region at the primary accreting pole dominated in the v -band [2]. since a two-pole accretion was determined from polarization of the optical emission for another high-state episode by [20], it suggests that the accretion modes are very unstable in am her. changes of the structure of the accretion region(s) are necessary to explain the variations of the optical and x-ray orbital modulation in the individual episodes of the high state in am her. the accretion modes (singlepole accretion, two-pole accretion) vary with time; they may be characteristic for a given high-state episode, but not for another one. since the properties of the accretion region at the primary pole largely differ from those at the secondary pole, these changes of the accretion modes contribute to the complicated long-term activity. the transition from the high to the low state is accompanied by dramatic changes of the size and structure of the cyclotron-emitting accretion region in am her (fig. 4). a decrease of the dimensions of the accretion region(s) during the transition emerges from these facts: the rises and declines between the extremes of brightness (especially of the branches of the primary minimum) become progressively steeper, the amplitude 83 v. šimon, a. henden of the modulation increases, the eclipse of the primary pole becomes very prominent. during the final phase of the high-state episode, the properties of the cyclotronemitting region do not scale exactly with the optical luminosity. am her appears to have the primary accreting region close to the position which is close to the grazing eclipse; a slight increase of its vertical and/or horizontal dimensions (and variable position) can give rise to a highly variable orbital modulation. this is visible not only for the site of the cyclotron emission (fig. 4), but also for the site of the bremsstrahlung emission which can increase so much that it even becomes temporarily uneclipsed in some high states [16]. figure 4: dramatic changes of the orbital modulation during transition from the high to the low state of am her (ccd v -band aavso data). time in jd– 2 400 000 is given in the legend. the orbital ephemeris of [5] is used. 5 conclusions in am her, the optical emission displays cycles in the transitions between the high and low states, which are detectable only for the limited time segments. in addition to the intermittent cycle of the transitions, accumulation of episodes of the low states in clusters produces an additional, long cycle of activity. a single isolated short episode of the low state does not suggest a break of this cycle. this has a consequence for the structures and their lifetimes on the donor and/or the changes of the configuration of the magnetic field in the whole system. the configuration of the accreting matter influences the emission of am her in various spectral bands. the accretion modes (single-pole accretion, two-pole accretion) vary with time (from one high-state episode to another). also the changes of the accretion region(s) and/or accretion modes are necessary to explain the relation between the optical and x-ray emission in the high states. acknowledgement this study was supported by grants 13-33324s and 13394643 provided by the grant agency of the czech republic. this research has made use of the observations provided by the asm/rxte team, the swift/bat team, the aavso international database (massachusetts, usa) and the afoev database (strasbourg, france). we thank the variable star observers worldwide whose observations contributed to this analysis. we used the code developed by dr. g. foster and available at http:// www.aavso.org/winwwz. we thank prof. petr harmanec for providing us with the code hec13. the fortran source version, the compiled version and brief instructions on how to use the program can be obtained via http:// astro.troja.mff.cuni.cz/ftp/hec/hec13/. references [1] foster, g.: 1996, aj, 112, 1709 [2] gänsicke, b. t., et al.: 2001, a&a, 372, 557 [3] gänsicke, b. t., et al.: 2006, apj, 639, 1039 doi:10.1086/499358 [4] heise, j., et al.: 1985, a&a, 148, l14 [5] heise, j., verbunt, f.: 1988, a&a, 189, 112 [6] henden, a.: 2011, aavso database [7] hessman, f. v., et al.: 2000, a&a, 361, 952 [8] hudec, r., meinunger, l.: 1976, ibvs, 1184, 1 [9] kafka, s., et al.: 2008, apj, 688, 1302 doi:10.1086/592186 [10] kafka, s., hoard, d. w.: 2009, pasp, 121, 1352 doi:10.1086/648579 [11] kalomeni, b.: 2012, mnras, 422, 1601 doi:10.1111/j.1365-2966.2012.20736.x [12] levine, a. m., et al.: 1996, apj, 469, l33 [13] livio, m., pringle, j. e.: 1994, apj, 427, 956 doi:10.1086/174202 [14] matt, g., et al.: 2000, a&a, 358, 177 84 http://dx.doi.org/10.1086/499358 http://dx.doi.org/10.1086/592186 http://dx.doi.org/10.1086/648579 http://dx.doi.org/10.1111/j.1365-2966.2012.20736.x http://dx.doi.org/10.1086/174202 activity of the polar am her (rx j1816.2+4952): a short review [15] michaut, c., et al.: 2012, mmsai, 83, 665 [16] priedhorsky, w., et al.: 1987, a&a, 173, 95 [17] šimon, v.: 2011, newa, 16, 405 doi:10.1016/j.newast.2011.03.001 [18] scharlemann, e. t.: 1982, apj, 253, 298 doi:10.1017/cbo9780511586491 [19] warner, b.: 1995, cataclysmic variable stars, cambridge univ. press, cambridge [20] wickramasinghe, d. t., et al.: 1991, mnras, 251, 28 [21] wu, k., kiss, l. l.: 2008, a&a, 481, 433 85 http://dx.doi.org/10.1016/j.newast.2011.03.001 http://dx.doi.org/10.1017/cbo9780511586491 introduction long-term activity in the optical band optical and hard x-ray intensities in the high states role of the accretion modes conclusions 107 acta polytechnica ctu proceedings 2(1): 107–110, 2015 107 doi: 10.14311/app.2015.02.0107 broad-band variability in accreting compact objects s. scaringi1,2 1instituut voor sterrenkunde, k.u. leuven, celesijlaan 200d, leuven, belgium 2max planck institute fur extraterrestriche physik, d-85748 garching, germany corresponding author: simo@mpe.mpg.de abstract cataclysmic variable stars are in many ways similar to x-ray binaries. both types of systems possess an accretion disk, which in most cases can reach the surface (or event horizon) of the central compact object. the main difference is that the embedded gravitational potential well in x-ray binaries is much deeper than those found in cataclysmic variables. as a result, x-ray binaries emit most of their radiation at x-ray wavelengths, as opposed to cataclysmic variables which emit mostly at optical/ultraviolet wavelengths. both types of systems display aperiodic broad-band variability which can be associated to the accretion disk. here, the properties of the observed x-ray variability in xrbs are compared to those observed at optical wavelengths in cvs. in most cases the variability properties of both types of systems are qualitatively similar once the relevant timescales associated with the inner accretion disk regions have been taken into account. the similarities include the observed power spectral density shapes, the rms-flux relation as well as fourier-dependant time lags. here a brief overview on these similarities is given, placing them in the context of the fluctuating accretion disk model which seeks to reproduce the observed variability. keywords: cataclysmic variables x-ray binaries optical timing photometry individual: mv lyrae individual: lu cam. 1 introduction cataclysmic variables (cvs) are close interacting binary systems where a late-type star transfers material to a white dwarf (wd) companion via roche lobe overflow. with an orbital period ranging from hours to minutes, the transferred material from the secondary star forms an accretion disc surrounding the wd. as angular momentum is transported outwards in the disc, material will approach the innermost regions close to the wd in the absence of strong magnetic fields, and eventually accrete on to the compact object. x-ray binaries (xrbs) are also compact interacting binaries which are similar to cvs in many ways, but where the accreting compact object is either a black hole (bh) or a neutron star (ns). both cvs and xrbs, as well as active galactic nuclei (agn; accreting supermassive bhs), have been shown to display strong aperiodic variability on a broad range of time-scales as well as in different wavelength ranges. xrbs have shown variability ranging from milliseconds to hours, whilst for cvs this ranges from seconds to days. this difference can be mainly attributed to the fact that the innermost edges of the accretion discs in cvs sit at a few thousand gravitational radii, whilst for xrbs it can reach down to just a few gravitational radii. the fact that material can get deeper within the gravitational potential of xrbs, as compared to cvs, also explains why they are more luminous and emit predominantly in x-rays, compared to cvs, which emit predominantly at optical/uv wavelengths. aperiodic broad-band variability (also referred to as flickering) has extensively been studied in x-rays for xrbs over several decades in temporal frequency (see for example terrell 1972; van der klis 1995; belloni et al. 2000; homan et al. 2001; belloni, psaltis & van der klis 2002). as cvs emit mostly at optical/uv wavelengths, timing studies of these objects had to rely on optical observing campaigns from earth, which are inevitably hindered by large interruptions, poor cadence, and in many cases poor signal-to-noise ratios. furthermore, the key time-scales to probe in cvs are much longer than in xrbs, requiring long, uninterrupted observations. recently, cv timing studies have been facilitated thanks to the advent of the kepler satellite (gilliland et al. 2010; jenkins et al. 2010), which is able to provide long, uninterrupted and high-precision light curves in the optical light from space. thanks to these capabilities it is now possible to probe over four orders of magnitude in temporal frequency in cvs. more importantly, it is now possible to compare the aperiodic variability properties observed in xrbs to those 107 http://dx.doi.org/10.14311/app.2015.02.0107 s. scaringi observed in cvs after taking into account the relevant timescale and wavelength “translations”. 2 broad-band aperiodic variability in accreting white dwarfs one important discovery relating the flickering properties observed in cvs to those observed in xrbs has come from studying the kepler lightcurve of the nova-like cv mv lyrae displaying typical flickering behaviour as observed in other cvs. specifically, scaringi et al. 2012a have reported the discovery of the so-called rms(root-mean-square)-flux relation within the kepler lightcurve of mv lyrae, where the rms-variability power linearly correlates with the mean cont rate of the source. this relation has been previously observed in a number of xrbs and agn (uttley & mchardy 2001; uttley et al. 2005), and, together with the observed log-normal flux distributions, rules out simple additive processes as the source of flicker noise (e.g. superposition of many independent shots), and instead strongly favours multiplicative processes (e.g. coupling of masstransfer variations travelling from the outer to inner disc for the latter) as the source of variability. more importantly, the fact that very similar rms-flux relations are found within all different types of accreting compact objects (bhs, nss, wds) on all scales strongly suggests that the driving mechanism responsible for the observed aperiodic variability is the same in all systems, irrespective of mass, size or type. additionally to displaying the rms-flux relation, scaringi et al. 2012b have reported that the power spectral density (psd) of mv lyrae is also qualitatively similar to those observed at x-ray wavelengths in xrbs. specifically, all psds display single or multiple quasi-periodic osscillations (qpos) as well as a high frequency break. both the psds of xrbs and cvs can be qualitatively modelled with a combination of lorentzian shaped functions, with the main difference arising from the characteristic frequencies involved. for example, xrbs display aperiodic variability on a wide range of timescales with high-frequency breaks at ≈ 100 − 101 hz (belloni et al. 2005), whilst cvs display very similar psd, but scaled to lower frequencies such that the high frequency break occurs at ≈ 10−3 hz. the difference between the psds in xrbs and cvs can be mainly attributed to the fact that material within the accretion disk can fall deeper within the embedded gravitational potential well of xrbs as opposed to cvs. this will result in xrbs displaying variability on shorter timescales than in cvs, and furthermore, will result in xrbs emitting most of their radiation at x-ray wavelengths as opposed to cvs emitting mostly at optical/uv wavelengths. 3 fourier-dependent time-lags in cvs similarities between the flickering properties of xrbs and cvs are not only limited to single-band observations. it has been known for over a decade that xrbs display high levels of coherence between two simultaneously observed x-ray lightcurves in different energy bands (vaughan & nowak 1997; nowak et al. 1999). associated to this, fourier-dependent time-lags are also observed at x-ray wavelengths, where hard xray photons are delayed with respect to the soft photons, with larger delays at the lowest temporal frequencies (known as hard lags). a natural explanation for hard lags in xrbs comes from the fluctuating accretion disk model (see arevalo & uttley 2006 and references therin), where the coupling of fluctuations in the mass-transfer rate at different accretion disk radii is responsible for the observed variability. as matter propagates from the cooler outer regions of the disk towards the hotter inner ones, the observed variability in two different energy bands will be delayed with respect to each other as a result of the temperature gradient in the accretion disk and fluctuations moving inwards. using the ultracam instrument (dhillon et al. 2007) mounted on the 4.2 meter william herschel telescope, scaringi et al. 2013 have obtained simultaneous optical lightcurves in the u′, g′ and r′ bands on the two nova-like systems mv lyrae and lu cam. the results from this observing campaign show that both systems display soft lags (where blue photons are observed before the red ones) at the lowest observed frequencies, with larger lags at low frequencies. soft-lags have also been observed at x-ray wavelengths in xrbs, as well as agn, and have been explained as photoionising reflection of hard-x-ray photons originating from the inner regions of the disk (possibly the corona) onto the outer regions. in this case, the lag delay would be the lighttravel time from the emitting region to the reflecting region. the soft lags observed in scaringi et al. 2013 are however too large (on the order of ≈ 10 seconds for lu cam) to be explained by light-travel time delay. nevertheless, it might still be that photons from the hot inner regions are being reprocessed by the cooler outer disk, possibly on the thermal timescale. 4 a physical model for the flickering variability the observed flickering properties in xrbs have been modelled in the context of the fluctuating accretion disk model (see arevalo & uttley 2006 and references therin). ingram & done (2011,2012) have shown how this model can be successfully applied to reproduce the observed psd shapes in xrbs. in the most recent at108 broad-band variability in accreting compact objects tempt by ingram & van der klis (2013), the authors have demonstrated how the model can be implemented analytically (rather than numerically), thus reducing the computation time of specific models from hours to seconds. in scaringi 2013 a simplification of the ingram & van der klis (2013) model has been applied to the kepler lightcurve of the cv mv lyrae by removing general relativistic effects which are negligible in accreting wds. the model prescription employed allows to fit the high-frequency psd of mv lyrae by associating the observed flickering to the viscous timescale at specific disk radii. assuming the accretion disk extends all the way to the wd surface, the model allows to fit for the disk outer edge, as well as for α(h/r)2, where α is the viscosity parameter and h/r the disk scale height (shakura & sunyaev 1973). a qualitatively good fit to the data is achieved (with reduced χ2 ≈ 1.2). the model parameters resulting from the fit are displayed in table 1, and suggest that the observed high-frequency flickering is driven by a geometrically thick disk extending from r ≈ 0.12r� all the way to the wd surface. table 1: best-fit parameters with associated (1σ) errors for the fluctuating accretion disk model applied to the kepler lightcurve of mv lyrae (scaringi 2013). mw d(m�) ≡ 0.73 rin(r�) ≡ 0.0105 rout(r�) 0.117 +0.029 −0.020 γ 0.853+0.047−0.041 α(h/r)2 0.705+0.289−0.182 fvar 0.220 +0.001 −0.001 although the observed emission at optical wavelengths in cvs is though to originate from a cold, geometrically thin outer disk, the results from this analysis seem to suggest that the flickering is driven by a geometrically thick inner disk. it is thus possible that the geometrically thin outer disk is reprocessing photons from the geometrically thick inner one, allowing to explain the inferred results. similar geometric configurations have also been inferred in xrbs. both x-ray timing and spectral analysis suggest that geometrically thick disk exists close to the compact object (optically thin, referred to as the corona), possibly sandwiching the geometrically thin, optically thick disk. if a similar configuration is confirmed in cvs, then it is possible that the physics responsible for generating the inner geometrically thick disks in cvs might be the same as that in xrbs. 5 conclusion both cvs and xrbs, as well as agn, are observed to display aperiodic broad-band variability. this variability can be associated to the accretion disks in these systems. as the accretion disks in xrbs can fall much deeper within the embedded gravitational potential well as opposed to cvs, most of their emission will occur at x-ray wavelengths as opposed to optical/ultraviolet. furthermore, the timescales associated with the disk inner edge are a few orders of magnitude higher in temporal frequency than those of cvs. here, a brief overview of the broad-band variability properties observed in cvs has been presented. in particular, the aperiodic properties discussed (psd shapes, rms-flux relations and fourier-dependent time-lags) are in many ways similar to those observed at x-ray wavelengths in xrbs once the relevant temporal scaling has been taken into account. the fluctuating accretion disk model which seeks to explain the observed variability in xrbs has also been briefly discussed in the context of cvs. this model associates the observed variability to the viscous timescale at specific disk radii. in this respect, applying this model to the kepler lightcurve of the nova-like cv mv lyrae suggests the existence of a geometrically thick disk close to the wd responsible for driving the observed high-frequency flickering. acknowledgement the author wishes to acknowledge funding from the fwo pegasus marie-curie fellowship program, as well as useful and insightful discussions with christian knigge, elmar koerding, phil uttley and tom maccarone. references [1] terrell 1972: apj, 174, l35. [2] van der klis 1995: x-ray binaries. cambridge: cambridge univ. press; 1995. p. 252. [3] belloni et al. 2000: a&a, 355, 271. [4] homan et al. 2001: apjs, 132, 377. doi:10.1086/318954 [5] belloni et al. 2002: apj, 572, 392. [6] gilliand et al. 2010: apj, 713, l160. [7] jenkins et al. 2010: apj, 713, l87. doi:10.1088/2041-8205/713/2/l87 [8] scaringi et al. 2012a: mnras, 427, 3396. doi:10.1111/j.1365-2966.2012.22022.x [9] uttley & mchardy 2001: mnras, 323, l26. [10] uttley et al. 2005: mnras, 359, 345. doi:10.1111/j.1365-2966.2005.08886.x 109 http://dx.doi.org/10.1086/318954 http://dx.doi.org/10.1088/2041-8205/713/2/l87 http://dx.doi.org/10.1111/j.1365-2966.2012.22022.x http://dx.doi.org/10.1111/j.1365-2966.2005.08886.x s. scaringi [11] scaringi et al. 2012b: mnras, 421, 2854. doi:10.1111/j.1365-2966.2012.20512.x [12] belloni et al. 2005: a&a, 440, 207 [13] vaughan & nowak 1997: apj, 474, l43. [14] nowak et al. 1999: apj, 510, 874. [15] lyubarskii 1997: mnras, 292, 679. [16] kotov 2001: mnras, 327, 799. [17] arevalo & uttley 2006: mnras, 367, 801. [18] dhillon et al. 2007: mnras, 378, 825. [19] caastella et al. 2012a: mnras, 422, 2407. [20] cassatella et al. 2012b: mnras, 427, 2985. doi:10.1111/j.1365-2966.2012.22021.x [21] de marco et al. 2013: mnras, 431, 2441. doi:10.1093/mnras/stt339 [22] scaringi et al. 2013: mnras, 431, 2535. doi:10.1093/mnras/stt347 [23] ingram & done 2011: mnras, 415, 2323. [24] ingram & done 2012: mnras, 419, 2369 doi:10.1111/j.1365-2966.2011.19885.x [25] ingram & van der klis 2013: mnras, 434, 1476. doi:10.1093/mnras/stt1107 [26] scaringi 2013: mnras, 438, 1233 [27] shakura & sunyaev 1973: apj, 24, 337. [28] narayan & yi 1994: apj, 428, 13. [29] fender et al. 2004: mnras, 355, 1105. discussion linda schmidtobreick: for nls we see a difference in the spectra between low and high inclination systems. low inclination systems show balmer absorption, possibly from an inner optically thick accretion disk which might coincide with your geometrically thick inner disk. does one see a similar relation of the presence of high frequency flickering with inclination? simone scaringi: the current model has only been applied on one system at the moment (mv lyrae, low inclination). future applications of the model on other systems will be able to determine whether the high frequency flickering displays different properties as a function of inclination. solen balman: how do you get an optically thick disk out of α(h/r)2 = 0.7? if α = 0.1 the disk is hot and it is no longer optically thick! simone scaringi: the inference of α(h/r)2 = 0.7 from the modelling of the psd in mv lyrae does not suggest that the disk is optically thick. it might well be that the inner region of the disk inferred here are optically thin as suggested. raymundo baptista: is you α constant with radius? simone scaringi: in the current model prescription α is kept as a constant with radius. however, it is more physical and realistic to allow α to have a radial dependence. this will be explored with the current model in the future. 110 http://dx.doi.org/10.1111/j.1365-2966.2012.20512.x http://dx.doi.org/10.1111/j.1365-2966.2012.22021.x http://dx.doi.org/10.1093/mnras/stt339 http://dx.doi.org/10.1093/mnras/stt347 http://dx.doi.org/10.1111/j.1365-2966.2011.19885.x http://dx.doi.org/10.1093/mnras/stt1107 introduction broad-band aperiodic variability in accreting white dwarfs fourier-dependent time-lags in cvs a physical model for the flickering variability conclusion 283 acta polytechnica ctu proceedings 1(1): 283–287, 2014 283 doi: 10.14311/app.2014.01.0283 the hiscore project m. tluczykont2, m. brückner1, n. budnev3, o. chvalaev3, a. dyachok3, s. epimakhov2, o. gress3, d. hampf2, d. horns2, a. ivanova3, n. kalmykov4, e. konstantinov3, e. korosteleva4, v. kozin4, m. kunnas2, l. kuzmichev4, b. lubsandorzhiev5, r. mirgazov3, r. monkhoev3, r. nachtigall2, m. panasyuk4, a. pakhorukov3, v. poleschuk3, a. porelli6, v. prosin4, v. ptuskin7, g.i. rubtsov5, p.s. satunin5, yu. semeney3, d. spitschan2, a. skukhin4, l. sveshnikova4, m. tluczykont2, r. wischnewski6, a. zagorodnikov3, v. zirakashvili7 1institute for computer science, humboldt-university berlin, rudower chaussee 25, 12489 berlin, germany 2institute for experimental physics, university of hamburg, luruper chaussee 149, 22761 hamburg, germany 3institute of applied physics isu, irkutsk, russia 4skobeltsyn institute for nuclear physics, lomonosov moscow state university, 1 leninskie gory, 119991 moscow, russia 5institute for nuclear research of the russian academy of sciences 60th october anniversary st., 7a, 117312, moscow, russia 6desy, platanenallee 6, 15738 zeuthen, germany 7izmiran, moscow, russia corresponding author: martin.tluczykont@physik.uni-hamburg.de abstract a central question of astroparticle physics, the origin of cosmic rays, still remains unsolved. hiscore (hundred*i square-km cosmic origin explorer) is a concept for a large-area wide-angle non-imaging air shower detector, addressing this question by searching for cosmic ray pevatrons in the energy range from 10 tev to few pev and cosmic rays in the energy range above 100 tev. in the framework of the tunka-hiscore project, first prototypes have been deployed on the site of the tunka-133 experiment, where we plan to install an engineering array covering an area of the order of 1 km2. on the same site, also imaging and particle detectors are planned, potentially allowing a future hybrid detector system. here we present the hiscore detector principle, its potential for cosmic ray origin search and the status of ongoing activities in the framework of the tunka-hiscore experiment. keywords: cosmic rays gamma rays instrumentation pevatrons. 1 introduction ever since their discovery, cosmic rays have triggered numerous experiments aiming at answering the questions about cosmic ray composition, spectral distribution, and about their origin. ground-based experiments measure cosmic ray or gamma ray induced air showers via particle or radiation detectors on the ground. the hiscore (hundred*i square-km cosmic origin explorer) concept (tluczykont et al. 2011), is based on the non-imaging air cherenkov technique, sampling the cherenkov radiation emitted by charged air shower particles. planned as a distributed array of sensitive light collecting stations (0.5 m2), with a wide field of view (0.60-0.84 sr) and a very large instrumented area (up to 100 km2), hiscore will cover gamma ray energies above 10 tev and spectral and composition measurements of cosmic rays from 100 tev up to the eev range. the main goal of hiscore is to find the cosmic ray pevatrons, that accelerate cosmic rays up to pev energies, and that are expected to emit gamma rays with a hard spectrum up to several 100 tev. air shower measurements at these energies open up several other research possibilities, (tluczykont et al. 2011). in the present paper, we describe the hiscore detector concept, its potential for a search for cosmic ray pevatrons, and the status of tunka-hiscore, currently (april 2013) in its deployment phase on the site of the tunka-133 experiment (berezhnev et al. 2012). 2 detector concept the hiscore detector is conceived as a distributed array of non-imaging light collecting stations. the array geometry planned for the tunka-hiscore engineering 283 http://dx.doi.org/10.14311/app.2014.01.0283 martin tluczykont et al. array is shown in figure 1. 150 m 106 m figure 1: array layout planned for the tunkahiscore engineering array. in the standard layout (simulated configuration) four 8 inch pmts are used per station (red squares). additional stations with larger pmts (10 or 11 inch) are envisaged for the engineering array (blue squares). first results of tests with different array layouts have shown that a graded array with a dense core and decreasing density toward the array edge provides significantly better angular resolution and maximizes the effective area. at the heart of each station, four 8 inch photomultipliers (pmts) measure the cherenkov light from secondary air shower particles. each pmt is equipped with light concentrators (winston cones) with a half opening angle of 30◦, covering a field of view of 0.84 sr. at reconstruction, a quality acceptance cut of 25◦ results in an effective field of view of 0.6 sr. the total light collection area of one station is 0.5 m2. after preamplification, the pmt signals are clipped and summed before passing a comparator. this clipped sum trigger concept prevents false triggers caused by large signal fluctuations from night sky background light or afterpulses. the signal amplitudes are digitized in time using a readout system based on the drs 4 (domino ring sampler) chip, operated at a sampling frequency of 1 ghz and a depth of 1 µs. signal amplitude, timing, rise time, and width of the arrival time distribution are used for the reconstruction of the direction, energy and nature of the primary particle. for a good reconstruction quality, the relative time-synchronization is required to be better than 1 ns between all stations of the distributed array. during one year, the 0.6 sr viewcone of the detector covers a total area of π sr for more than 200 h observation time. while this area is much larger than areas typically covered in scans performed by imaging cherenkov telescope arrays, it is fixed by the location of the experiment. in order to overcome this deficit, we plan to reorient the detector stations after a few years of operation, tilting the detector axis in north-south direction, as illustrated in figure 2, extending the sky coverage and partly deepening the exposure. figure 2: schematical view of a tunka-hiscore detector station. the tilting mechanism will allow to access a larger fraction of the sky as compared to observations pointing at the zenith alone. detail (right): each of the four optical channels consists of a pmt with a light concentrator (winston cone). 3 sensitivity the sensitivity of the hiscore standard layout was computed on the basis of a full simulation (tluczykont et al. 2011) and reconstruction algorithms (hampf, tluczykont & horns 2013). we require a detection at the level of 5 standard deviations and at least 50 gamma rays, leading to the point-source survey sensitivity shown in figure 3. the blue area corresponds to the sensitivity of the hiscore non-imaging array in its standard configuration (8 inch pmts, 150 m interstation spacing). the upper edge of the sensitivity range was obtained using conservative assumptions, assuming a ratio α of the solid angles of signal to background regions of unity. the gamma ray survival probabilities are 0.68 for the angular point-source cut, and 0.6 for the hadron rejection cut. more optimistic assumptions were made for the lower bound of the dark shaded area. here, α << 1 and the angular cut was relaxed to an efficiency of 1 in the background free regime, above 2 pev. for direct comparison, the survey sensitivity of cta (dubus et al. 2013) is shown as a dashed area, and the lhaaso sensitivity adapted from cao et al. 2010 with an additional requirement of at least 50 gamma rays is shown as dash-dotted curve. since the gamma-hadron rejection is poor when using the nonimaging technique alone, the impact of an improvement of the hadron rejection efficiency is significant. for illustration, the cyan area shows the sensitivity range when including an improvement of the hadron rejection by a factor of 10. such an improvement is realistic when combinining the hiscore concept with imaging or muon detectors. imaging detectors have proven 284 the hiscore project their gamma-hadron separation power up to several 10s of tev. at higher energies with a rising average number of muons per air shower, muon detectors become more and more efficient for hadron rejection. hiscore will open up the pevatron energy range and allow detailed spectral measurements of the continuation of known gamma-ray sources in the multi-tev to pev energy regime. at these energies, the absorption of gamma rays (e+e− pair production) in the interstellar radiation field (isrf) around 100 tev and the cosmic microwave background (cmb) around 3 pev are significant. the level of gamma ray attenuation depends on the density of the radiation fields and the distance of the objects, and can be as low as few percent. moreover, the density of the cmb being known, the absorption expected at pev energies from the cmb might open up a new method for distance estimation using gamma ray observations. energy / tev -110 1 10 210 310 410 -1 s -2 in te g ra l f lu x / e rg c m -1410 -1310 -1210 -1110 -1010 hiscore cta survey lhaaso point-source, 1 year )νicecube milagro sources, 5 years ( kascade u.l. mgro j1908+06 hess j1908+06 hiscore figure 3: hiscore sensitivity as compared to data on mgro j1908+06 from milagro (abdo et al. 2007) and h.e.s.s. (aharonian et al. 2009), and the sensitivity of icecube to neutrinos from milagro sources (gonzalez-garcia et al. 2009). also given are an upper limit from kascade (antoni et al. 2004), and sensitivities for experiments cta (dubus et al. 2013) and lhaaso (cao et al. 2010). recently, the icecube collaboration reported the detection of 28 neutrinos (e.g. whitehorn 2013), in excess of the atmospheric and prompt neutrino expectations, but compatible with a pevatron like spectrum. these findings are a strong additional motivation for our search for galactic cosmic ray pevatrons. 4 status of tunka-hiscore 4.1 detector components different solutions are available for the individual detector components. the photomultipliers (pmts) are required to sustain a high level of noise and to operate at a gain level of 104 105. currently, 8 inch pmts fulfilling these requirements are available from emi, electrontubes and hamamatsu (including a modified 6-stage prototype). the winston cones are assembled similarly to the construction of a wine barrel in stripes along the optical axis, using reflective sheets of alanod 4300up material. in the currently operating prototypes (april 2013), standard tunka-133 preamplifiers are in use. an analog summator for the station trigger system (see section 2) was tested in laboratory and in the field. in its final implementation, the trigger board includes clipping and triggering. further three trigger solutions are available: fadc trigger onboard the tunka-133 readout, fpga based trigger on-board the white rabbit board, and amplitude discrimination on-board the drs 4 evaluation board. all three alternative triggers use the analog sum as input. two readout systems are in operation: the standard fadc readout (200 mhz) used by tunka-133, and the drs 4 evaluation board (1 ghz). 4.2 prototype results the first tunka-hiscore prototype station (2 pmt channels) was deployed in the tunka valley in april 2012. two additional detector stations were deployed in october 2012. joint operation with tunka since then provided first cross-calibration data and experience with components. the tunka-133 fadc and the drs 4 evaluation board readout were tested in parallel in spring 2013. runs with varying thresholds were performed in order to measure the integral trigger rate as a function of the trigger threshold (epimakhov et al. 2013). as can be seen in figure 4, the expectations for the cosmic ray branch and the noise wall can be reproduced. threshold [mv] 210 t ri g g e r ra te [ h z ] 1 10 210 310 410 510 610 710 figure 4: the trigger rate (station 4) as a function of the trigger threshold, measured with the drs 4 evaluation board in spring 2013 (epimakhov et al. 2013). red squares: dead-time corrected data, red dashed line: cosmic ray triggers, black points: estimation of nsb triggers from baseline fluctuations, black dashed curve: the nsb noise wall. 285 martin tluczykont et al. similar measurements using the fadc readout yield consistent results (wischnewski et al. 2013). an effective quantum efficiency of 0.06 could be derived from the latter measurement. this translates into a threshold for cherenkov light detection of 0.6 photons/m2, which corresponds to an energy threshold of slightly less than 100 tev when using the fadc trigger (25 ns discriminator response). the final system will have a faster response. we also plan to use faster preamplifiers, larger pmts, and geo-magnetic shielding for the pmts, ultimately lowering the energy threshold of the hiscore stations to the 10–30 tev energy range. tests of the white rabbit (wr) system show that the time precision and stability at the 200 picos level measured in the laboratory can be reproduced at the tunka site (brückner et al. 2013). the observed time difference distribution of wr-tagged events is consistent with the observation of air showers with the given viewing angle of the tunka-hiscore stations (wischnewski et al. 2013). 5 conclusion and outlook using the new large-area wide-angle non-imaging hiscore detector concept, we plan to access the pevatron gamma ray energy range (10 tev – several 100 tev) and also address cosmic ray physics around the knee region. the tunka-hiscore experiment has started the deployment of detector stations in the tunka valley in siberia. 3 prototype stations are in operation, and further 6 stations will be deployed by the end of 2013. the construction of an engineering array that will cover an area of 1 km2 is planned for the near future. different alternatives for the station design (8 inch or 10 inch pmts) and for the array layout (graded array) are considered. a combination with small imaging telescopes (2-3 m2 mirrors, 10◦ field of view) is planned to complement the tunka-hiscore array. furthermore, the installation of muon detectors and fluorescence telescopes at the tunka site is considered for the future, improving gamma hadron separation and cosmic ray composition, and leading to a multi component extended air shower array in the tunka valley in siberia. acknowledgement we acknowledge the support of the russian federation ministry of education and science (g/c 14.518.11.7046, agreements n 14.b25.31.0010, n 14.b37.21.0785, and n 14.37.21.1294, g/c 14.740.11.0890, p681, contract n 14.518.11.7046), the russian foundation for basic research (grants 11-0200409, 13-02-00214, 13-02-12095, 13-02-10001, 12-0210001, 12-02-91323), the president of the russian federation (grant mk-1170.2013.2), the helmholtz association (grant hrjrg-303), and the deutsche forschungsgemeinschaft (grant tl 51-3). we thank gavin rowell for valuable discussions. references [1] abdo, a. a., allen, b., berley, d., et al.: 2007, apj l., 664, l91 doi:10.1086/520717 [2] aharonian, f., akhperjanian, a. g., anton, g., et al.: 2009, a&a, 499, 723 [3] antoni, t., apel, w. d., badea, a. f., et al.: 2004, apj, 608, 865 doi:10.1086/420736 [4] berezhnev, s. f., et al. (tunka collaboration): 2012, nima, 692, 98 [5] brückner, m., et al.: 2013, to appear in proc. of icrc, rio de janeiro, brazil, 1158 [6] cao, z., bi, x. j., cao, z., et al.: 2010, 38th cospar scientific assembly, 38, 2322 [7] dubus, g., contreras, j. l., funk, s., et al.: 2013, aph, 43, 317 [8] epimakhov, s., brückner, m., budnev, n. et al.: 2013, to appear in proc. of icrc, rio de janeiro, brazil [9] gonzalez-garcia, m. c., halzen, f., & mohapatra, s.: 2009, aph, 31, 437 [10] hampf, d., tluczykont, m., & horns, d.: 2013, nimpa, 712, 137 doi:10.1016/j.nima.2013.02.016 [11] moskalenko, i.v., porter, t.a., strong, a.w.: 2006, apjl, 640 l155 doi:10.1086/503524 [12] tluczykont, m., hampf, d., horns, d., et al.: 2011, adv. sp. res., 48, 1935 doi:10.1016/j.asr.2011.08.004 [13] whitehorn, n. h.: 2013, fermilab astro seminar [14] wischnewski, r., berezhnev, s., brückner, m., et al.: 2013, to appear in proc. of icrc, rio de janeiro, brazil, 1164 discussion (from memory) questioner: at these high energies, gamma rays are absorbed in low energy radiation fields. how can you expect sources in the 100 tev energy range? martin tluczykont: the strength of the attenuation depends on the location of the object within our galaxy. an object located at the far side of 286 http://dx.doi.org/10.1086/520717 http://dx.doi.org/10.1086/420736 http://dx.doi.org/10.1016/j.nima.2013.02.016 http://dx.doi.org/10.1086/503524 http://dx.doi.org/10.1016/j.asr.2011.08.004 the hiscore project the galaxy behind the galactic center was shown (moskalenko et al. 2006) to be absorbed at the 50 % level. any other object located closer to the sun, or off-center the galactic disk, will be less absorbed, down to only few %. 287 introduction detector concept sensitivity status of tunka-hiscore detector components prototype results conclusion and outlook 307 acta polytechnica ctu proceedings 1(1): 307–310, 2014 307 doi: 10.14311/app.2014.01.0307 recent results from the safir project1 m. sánchez-portal1,2, m. castillo-fraile1,2, c. ramos almeida3, p. esquej4,5, a. alonso-herrero5, a. m. pérez garćıa3, j. acosta-pulido3, b. altieri1, a. bongiovanni3, j. m. castro cerón1, j. cepa3, d. coia1, l. conversi1, j. fritz6, j. i. gonzález-serrano5, e. hatziminaoglou7, m. pović8, j. m. rodŕıguez espinosa3, i. valtchanov1 1european space astronomy centre (esac)/esa, madrid, spain 2isdefe, madrid, spain 3instituto de astrof́ısica de canarias, la laguna, tenerife, spain 4centro de astrobioloǵıa, inta-csic, madrid, spain 5instituto de f́ısica de cantabria, csic-uc, santander, spain 6sterrenkundig observatorium, universeit gent, belgium 7european southern observatory, garching bei münchen, germany 8instituto de astrof́ısica de andalućıa, granada, spain corresponding author: miguel.sanchez@sciops.esa.int abstract the “seyfert and star formation activitiy in the far-infrared” (safir) project is aimed at studying the physical nature of the nuclear ir emission and star formation properties of a small sample of nearby seyfert galaxies observed with the pacs and spire instruments on board the herschel space observatory. in this paper, we review the achieved results, that reveal the importance of the far-ir range to improve the quality and reliability of the estimates of basic agn torus parameters, and describe some preliminary outcome from the on-going work on the dust properties of resolved agn host galaxies. keywords: agn seyfert sed ir. 1 introduction coeval agn and starburst phenomena can be assessed by means of the analysis of dust in the infrared (ir) domain. in this range, in particular in the mid-ir (mir) and far-ir (fir), dust contributes to most of the thermal emission. the unified model considers a central agn engine and a broad-line region (blr) obscured by a thick dust torus. the dust grains re-radiate in the ir the absorbed uv/optical photons. as it has been well characterised by existing facilities (eg. spitzer, trecs), the dusty torus emission peaks in the mir (730 µm) and it extents to the fir, where the contribution related to the star formation (sf) becomes dominant. thus, agreeing to [8], the sed of seyfert galaxies in the mir and fir range can be solely explained by the dust thermal re-radiation of higher energy photons. therefore, dust thermal emission should be made-up of three different contributions: (a) warm dust heated by the agn (120-170 k); (b) cold dust heated by the star formation (40-70 k, and (c) very cold dust heated by the general interstellar radiation field (15-25 k). until now it has been poorly constrained due to the limited spatial resolution and spectral coverage of the existing facilities. the herschel observatory [5] provides new performances and capabilities to study the emission of nearby galaxies in fir ans sub-mm regions: the pacs [9] photometer allows to image in the 70, 100 and 160 µm with unprecedented spatial resolution (5.5 arcsec at 70 µm) and the spire [3] photometer permits to image in the 250, 350, 500 µm bands, a formerly unexplored region, at a relatively high spatial resolution. these instruments provide, on the one hand the characterisation of the agn sed minimizing the contamination by the host galaxy (pacs) and on the other, the assessment of the cold and very cold dust components (spire) both across the host galaxy and the nuclear and circum-nuclear regions. pacs and spire data can 1herschel is an esa space observatory with science instruments provided by european-led principal investigator consortia and with important participation from nasa. 307 http://dx.doi.org/10.14311/app.2014.01.0307 miguel sánchez-portal et al. be used to fit seds sampling both the emission peak and the rayleigh-jeans tail of the thermal emission of the cold and very cold dust components. the fitted seds can be used to derive dust masses and temperatures and the star formation rate (sfr). the relatively high spatial resolution of these instruments makes possible to map the different emission regions (e.g. nucleus, arms, inter-arm region). in this context, two foci of interest drive the safir study: (a) the dusty torus: current models consider either smooth or clumpy dust distributions. any agn ir model should consider three constituents when applied to describe the actual sed of seyfert galaxies: agn, starburst and host galaxy. the starburst contribution can be constrained by the use of fir data. in addition, both the torus and starburst emission overlap smoothly in the fir. therefore, the use of fir data with high spatial resolution and wide spectral range coverage is fundamental to discriminate the modelled torus characteristics. (b) the nuclear activity and star formation coexistence: the interrelationship between accretion onto massive black holes and the star formation is a topic fundamental to understand the formation and evolution of galaxies. fir data from herschel pacs and spire allow to characterise and map the dust distribution and temperature and the star formation activity across galaxies hosting agns. it can contribute to constraint the current agn formation and evolution models. 2 sample of galaxies and technical implementation the safir collection of nearby galaxies is constituted by 18 seyfert galaxies sampled to represent different nuclear clases (seyfert 1.x & seyfert 2). only objects with available high-resolution mir data (groundbased or spitzer) were selected. in addition, all the objects have available optical, nir, x-radio and radio data. this allows to construct a complete multiwavelength sed for all the objects of the sample. ten galaxies of the sample are barred spiral/lenticulars and five are peculiar/interacting systems. four objects are confirmed luminous or ultra-luminous ir galaxies (lirg/ulirg). the observations were performed in the pacs and spire scan map modes adjusting the mapped areas to fit the host galaxy and a background region within the surveyed area. the achieved 1σ sensitivities were approximately, 3.6, 3.9 and 3.9 mjy/beam for pacs at 70,100 and 160 µm and 5.5, 7.6 and 6.4 mjy/beam for spire at 250, 350 and 500 µm. with these sensitivities it was possible to map both the nuclear and circum-nuclear regions and also large areas within the galaxy disks. 3 results this section presents some already published results from the safir project [10, 2, 1] for three objects of the sample. 3.1 ngc 3081 this galaxy was studied in the context of the safir project combining pacs/spire data with groundbased high-resolution nir/mir data [10]. this object is an early-type barred spiral ((r)sab0/a(r)). it comprises a series of well defined nested star-forming annular-like features: nuclear (r1, 2.3 kpc), inner (r2, 11 kpc) and outer (26.9 kpc) rings. the inner ring (r2) is evidently resolved in the images up to 250 µm. the nuclear sed was fitted combining unresolved fir fluxes (r≤1.7 kpc) together with integrated nir and mir data. a clumpy model was applied to simulate the nuclear torus emission [4] to assess how the the torus parameters have to be modified, in particular the torus size to account for the fir emission. as a result, it was obtained that the torus outer radius must be notably increased: ro = 4 +2 −1 pc vs. ro = 0.7±0.3 pc obtained using only nir and mir data [11]. also the radial distribution of clouds (defined by the power-law index of the radial density profile q) flattens when the fir data are included in the simulation: (q = 0.2 vs q = 2.3). other model parameters (width of the angular distribution, inclination angle, optical depth, number of clouds) are in agreement with those obtained without fir data. at larger scales (1.7 kpc≤r≤5.4 kpc), the fir emission is well characterised by cold dust thermal emission at t = 28±1 k (assuming a grey blackbody with emissivity β = 2) likely heated by young stars in r1. the fir emission of the outer part of the galaxy can be reproduced with very cold dust (t = 19±3 k) heated by the interstellar radiation field. 3.2 mrk 938 this galaxy contains a seyfert 2 agn and presents a significant starbust activity. this object is a morphologically peculiar galaxy that has been proposed to be the remnant of a gas-rich merging of two unequal mass galaxies [12]. it is classified as lirg due to its large ir luminosity. a multi-wavelength study was performed for this object combining x-ray, nir, mir and pacs/spire fir data in the context of the safir project [2] in order to characterise the origin and nature of its strong emission in the ir range. the agn bolometric contribution to the mir and the total ir luminosity is small [lbol(agn)/lir∼0.02] as observed in the component decomposition of the mir spitzer/irs spectrum, which is in agreement with previous estimations. the mips 24 µm and pacs 70 µm images indi308 recent results from the safir project cate that the major part of the star formation activity is concentrated in a compact obscured region of ≤2 kpc. fir data have been used to constraint the cold dust emission with unprecedented accuracy. in order to derive the dust properties, the integrated ir sed has been fitted. it has been found that the mir to fir spectrum can be properly modelled by a two-component sed: two modified blackbodies with fixed emissivity (β = 2) and temperatures tw = 67 k and tc = 35 k for the warm and cold dust components, respectively. in addition, a single blackbody component sed was fitted to the fir-only spectrum with a modified blackbody of β = 2 and t = 36.5 k. this value has been used along with the spire flux at 250 µm to derive the dust mass using eq. 2 considering an absorption coefficient κ250µm = 4.99 cm 2g−1. the value obtained for the dust mass, mdust = 3×107 m�, is consistent with those derived for local ulirgs and other ir-bright galaxies. 3.3 ngc 1365 ngc 1365 (fig. 1, left) is a supergiant barred spiral galaxy (sb(s)b). it is a nearby (18.6 mpc) lirg harboring a seyfert 1.5 type nucleus. the inner linblad resonance (ilr) region of the galaxy contains a powerful nuclear starburst ring with an approximate diameter of 2 kpc. we have probed the nuclear and circum-nuclear activity of this galaxy in the ir [1]. the strong star formation activity in the ring is resolved by the herschel/pacs imaging data that shows some substructures (super star clusters), as well as by the spitzer 24 µm continuum emission, [ne ii]12.81 µm line emission, and 6.2 and 11.3 µm pah emission. the active galactic nucleus (agn) is the brightest source in the central region up to λ∼24 µm, but it becomes increasingly fainter in the fir when compared to the emission originating in the ir clusters located in the ring. we modelled the agn unresolved ir emission with a clumpy torus model and estimated that the agn contributes only in a small fraction (∼5%) of the ir emission produced in the inner ∼5 kpc. the estimated torus size is ∼5 pc. we fitted the non-agn 24–500 µm sed of the region within the ilr and found that the dust temperature and mass are similar to those of other nuclear and circum-nuclear starburst regions. finally, the comparison of the ir-derived sfr with that obtained from hα observations indicates that ∼85% of the ongoing star formation within the ilr is taking place in dust–obscured regions. 4 dust properties of resolved agn host galaxies a study of the dust properties of spatially well-resolved agn hosts has been started (sánchez-portal, castillofraile et al. in preparation). within the safir sample, four galaxies (ngc 1365, ngc 4258, ngc 1566 and ngc 5728) have an apparent size large enough to allow a detailed analysis of the spatial dust properties, notably its temperature and mass that can be directly compared with the star formation characteristics. for these objects, the spatial resolution of the observations is being exploited to produce maps of the dust mass, temperature, and sfr. in this section some examples of the activities currently on-going are shown. assuming an optically thin emission, the flux density can be expressed as fν ∝ νβb(ν,tdust) where β is the dust emissivity. as already stated, several dust components with different temperature should be generally considered, so the flux density can be expressed as: fν(λ) = n∑ i=1 ni λβ+3(ehc/λkti − 1) (1) figure 1: temperature maps of ngc 1365 (left) and ngc 1566 (right). average temperatures range from ∼ 17–18 k in the inter-arm regions to t ∼23–24 k in the bright spots within the spiral arms. the highest average dust temperatures are observed in the central region of ngc 1365 with t ∼26 k. where ni are the normalization constants and ti are the temperatures of the different components. the procedure devised to generate temperature maps includes the following steps: after a standard reduction procedure (see [1] for a description), the pacs 70, 100 and 160 µm and spire 250 and 350 µm maps have been convolved to the resolution of the spire 500 µm images and resampled to the largest pixel size (that of the spire 500 µm maps, set to 14 arcsec). the images have been spatially registered and used as input to an idl procedure that performs a least-squares fit to either one or two dust components (n = 1 or 2) at each pixel in order to cope with the cold or/and very cold dust components. in the maps shown in fig. 1 we have used a single temperature component with a fixed emissivity β = 2 to create the temperature maps of ngc 1365 (top) and 309 miguel sánchez-portal et al. ngc 1566 (bottom). the latter is a bright (lc ii-iii), nearby (11.83 mpc) sab(s)bc spiral galaxy harboring a seyfert type 1.5 nucleus. figure 2: sfr map of ngc 1365, obtained from the grey body ir luminosity integrated between 8 and 1000µm. the temperature maps generated closely follow the topology of the star formation regions, with the highest temperatures corresponding to areas of high sf activity, as observed by comparison with the morphology of 70 µm and optical hα images. in fact, we have created sfr maps by integrating the grey body sed at the best-fit temperature and applying standard scaling relations [7]. in fig. 2 we show the sfr map of ngc 1365. there is an excellent agreement with the structures revealed by the dust temperature map and the sfr density. in agreement with [1], it is observed that the most intense star formation is taking place in the circum-nuclear region (within the ilr). outstanding formation rate is also taking place in the spiral arms. figure 3: ngc 4258 temperature (left) and dust mass (right). the spatial distribution of dust mass (projected dust density) can be obtained from the temperature maps, using the expression: mdust = d2lfν κνbν (tdust) (2) adapted from [6], where dl is the luminosity distance and fν is extracted from the spire flux map at 250 µm assuming an absorption coefficient κ250µm = 4.99 cm 2g−1. in fig. 3 we present the dust temperature and mass maps of ngc 4258, a bright (lc ii-iii), nearby (7.44 mpc) sab(s)bc spiral galaxy hosting a liner/seyfert 1.9 nucleus. the pixel scale is 0.36 kpc−2. 5 conclusions the high spatial resolution herschel pacs & spire observations are demonstrating the importance of the fir to improve the quality and reliability of the agn torus fits. parameters as important as the torus radius and cloud radial distribution have a strong dependency of this spectral range. moreover, the fir data are crucial to characterise the starburst contribution and to constrain the dust properties. the quality of the herschel data is allowing us to study the spatial distribution of dust within the galaxies, thus permitting to characterize the variation of dust properties (temperature, dust mass) and sfr with the nuclear distance. acknowledgments we would like to acknowledge the herschel project scientist, göran pilbratt, for making possible the implementation of this project kindly providing the required guaranteed time from the ps budget. references [1] alonso-herrero, a., sánchez-portal, m., ramos almeida, c., et al. 2012, mnras, 425, 311 doi:10.1111/j.1365-2966.2012.21464.x [2] esquej, p., alonso-herrero, a., pérez-garćıa, a. m., et al. 2012, mnras, 423, 185 doi:10.1111/j.1365-2966.2012.20779.x [3] griffin, m.j., abergel, a., abreu, a. et al. 2010, a&a, 518, l3 [4] nenkova m. et al., 2008, apj, 685, 147 doi:10.1086/590482 [5] pilbratt, g.l., riedinger, j.r., passvogel, t. et al. 2010, a&a, 518, l1 [6] hildebrand r. h., 1983, q. j. r. astron. soc., 24, 267 [7] kennicutt, jr., r. c., araa, 1998, 36, 189 doi:10.1146/annurev.astro.36.1.189 [8] pérez garćıa a. m., rodŕıguez espinosa j. m., 2001, apj, 557, 39 doi:10.1086/321675 [9] poglitsch, a., waelkens, c., geis, n. et al. 2010, a&a, 518, l2 [10] ramos almeida, c., sánchez-portal, m., pérez garćıa, a. m., et al. 2011, mnras, 417, l46 doi:10.1111/j.1745-3933.2011.01117.x [11] ramos almeida c. et al., 2011, apj, 731, 92 doi:10.1088/0004-637x/731/2/92 [12] schweizer f., seitzer p., 2007, aj, 133, 2132 310 http://dx.doi.org/10.1111/j.1365-2966.2012.21464.x http://dx.doi.org/10.1111/j.1365-2966.2012.20779.x http://dx.doi.org/10.1086/590482 http://dx.doi.org/10.1146/annurev.astro.36.1.189 http://dx.doi.org/10.1086/321675 http://dx.doi.org/10.1111/j.1745-3933.2011.01117.x http://dx.doi.org/10.1088/0004-637x/731/2/92 introduction sample of galaxies and technical implementation results ngc 3081 mrk 938 ngc 1365 dust properties of resolved agn host galaxies conclusions 205 acta polytechnica ctu proceedings 1(1): 205–209, 2014 205 doi: 10.14311/app.2014.01.0205 expected hard x-ray and soft gamma-ray from supernovae keiichi maeda1,2, yukikatsu terada3, aya bamba4 1department of astronomy, kyoto university, japan 2kavli institute for the physics and mathematics of the universe (wpi), university of tokyo, japan 3department of physics, saitama university, japan 4department of physics and mathematics, college of science and engineering, aoyama gakuin university, japan corresponding author: keiichi.maeda@kusastro.kyoto-u.ac.jp abstract high energy emissions from supernovae (sne), originated from newly formed radioactive species, provide direct evidence of nucleosynthesis at sn explosions. however, observational difficulties in the mev range have so far allowed the signal detected only from the extremely nearby core-collapse sn 1987a. no solid detection has been reported for thermonuclear sne ia, despite the importance of the direct confirmation of the formation of 56ni, which is believed to be a key ingredient in their nature as distance indicators. in this paper, we show that the new generation hard x-ray and soft γ-ray instruments, on board astro-h and nustar, are capable of detecting the signal, at least at a pace of once in a few years, opening up this new window for studying sn explosion and nucleosynthesis. keywords: nuclear reaction nucleosynthesis abundances supernovae: general. 1 introduction supernova (sn) explosions trigger (or are triggered by) explosive nucleosynthesis, and they are believed to be main production sites of heavy elements in the universe. the resulting yields are sensitive to explosion mechanism(s), and thus studying nucleosynthesis products is important to uncover the still-debated explosion mechanism. especially important is the production of 56ni – this is the origin of fe (as a result of the radioactive decay chain 56ni → 56co → 56fe), and the decay is believed to provide a source of emissions from (many classes of) sne through thermalization of emitted γ-rays and positrons. in type ia supernovae (sne ia), about half of an exploding white dwarf in mass is processed into 56ni, supporting their huge luminosities as distance indicators. however, the most direct evidence in this scenario is still missing – there has been no solid detection of the decay γ-rays from sne except for sn 1987a (e.g., dotani et al., 1987; sunyaev et al., 1987). especially, no solid detection has been reported for sne ia (see milne et al., 2004 for a review). from a theoretical point of view, studying this high energy emission has been restricted to one-dimensional models (see milne et al., 2004, for a review) despite the importance of multi-dimensional structures of the explosion both in theory and observation (e.g., kasen et al., 2009; maeda et al., 2010a). most previous studies also focused on the emission in the mev range. in this paper, we present our radiation transfer simulations of the high energy emission based on the state-of-the-art sn ia explosion models. we extend our analysis to hard x-ray and soft γ-ray regimes, for which dramatic improvement is expected in the observational sensitivities thanks to new generation observatories like nustar (koglin et al., 2005) or astro-h (takahashi et al., 2010). we predict that these telescopes are capable of detecting the radioactive decay signals from sne ia, at a rate of once in a year or at least once in a few years. we also briefly comment on perspectives for core-collapse sne. 2 expected characteristics we performed radiation transfer simulations (maeda et al. 2012) based on a series of two-dimensional delayed detonation models by kasen et al. (2009). the delayed detonation model is among the most popular scenarios for sne ia, resulting from a near-center ignition of thermonuclear sparks within a chandrasekhar-mass white dwarf (khokhlov, 1991). conditions for the initial triggers have not been clarified from the first principle (e.g., seitenzahl et al., 2013), thus kasen et al. (2009) adopted various conditions (i.e., distribution of the sparks) and produced a range of the ejecta models. in this scenario, different initial conditions can be associated with observed diversities. figure 1 shows examples of the ejecta structure. the model dd2d asym 04 is for bright sne ia (resulting in ∼ 1m� of 56ni), while 205 http://dx.doi.org/10.14311/app.2014.01.0205 keiichi maeda, yukikatsu terada, aya bamba figure 1: examples of the delayed-detonation models (kasen et al., 2009; see also maeda et al., 2010b). the mass fractions of si (left) and 56ni (right) are shown, on a logarithmic scale. the axes are in 10, 000 km s−1. dd2d iso 04 is for fainter ones (∼ 0.4m� of 56ni). generally, this model sequence predicts more asymmetric structure for fainter sne ia (note that the initial condition of model dd2d asym 04 is indeed more asymmetric than dd2d iso 04, but the post-explosion ejecta are less asymmetric). examples of the synthetic spectra are shown in figure 2. the spectra are characterized by the decay lines, compton scattering continuum, and the low energy cut off by the photoelectric absorption. at 20 days, the decay lines from 56ni → 56co (e-folding time of ∼ 8.8 days) are more important than those from 56co → 56fe (∼ 113 days) as characterized by strong lines in the soft γ-ray range (e.g., the 158 kev line followed in strength by the 270 and 480 kev lines). later on, the strong lines are mostly in the mev range (i.e., the 847 kev line as the strongest) except for the annihilation line. the cut off energy becomes higher as time goes by due to the increasing contribution to the emission from the deeper part where the mean atomic number and photoelectric cross sections are larger. thus, overall the spectra evolve from soft to hard as time goes by. this indicates that follow-up of sne at relatively early phases is important in hard x-ray and soft γ-ray range. the classical 1d model w7 has lower energy cut off than the delayed detonation model sequence, since the w7 model has less extended explosive nucleosynthesis in the surface layer than the delayed detonation model. due to increasing transparency, the emission is sensitive to model variants (including the viewing direction) in the earlier phase, while the mass of 56ni plays a dominant role in the later phase. thus, observation at relatively early phases in (relatively) soft bands can provide unique diagnostics in clarifying the explosion mechanism(s). in the multi-d delayed detonation model sequence, a unique prediction is that faint sne ia should show larger variations in the high energy emission arising from various viewing directions than brighter ones. such prediction can be tested once there are at least a few samples of high energy emission detected for sne ia. another diagnostics using the ‘soft’ bands includes the surface composition analysis from the photoelectric absorption, e.g., the different behavior shown for the w7 model and delayed detonation models. 3 observational perspectives table 1 summarizes expected detectability of extragalactic sne ia at various band passes by a few current and future instruments. while detecting the mev lines from the 56co decay has been challenged in the past, even with spi on board integral and 106 s exposure (roques et al., 2003; see also isern et al., 2013), this is limited to extremely nearby sne ia up to 5 6 mpc (or 8 mpc for extremely bright sne ia). such nearby events are expected only once in a decade. this frustrating situation in the mev range will be improved only when the sensitivity is improved by an order of magnitude, hopefully by proposed new generation observatories like grips (greiner et al., 2012; see also summa et al., 2013). we propose that new generation hard-x and/or soft γ-ray instruments can change the situation. in hard xrays, nustar has been already launched. astro-h is scheduled for launch in 2015, which will be attached with hxi (hard x-ray) and sgd (soft γ-ray). these instruments are expected to have 106 s exposure sensitivities sufficient to reach sne ia at 15 (conservative estimate) or 25 mpc (optimistic estimate) (tab. 1). the ‘line detection’ is more challenging, and we estimate the distance for the 5σ detection of the 158 kev line is ∼ 10 − 15 mpc for model dd2d asym 04 and ∼ 3 − 5 mpc for the other two models shown in table 1. figure 3 shows simulations for expected signals from sne ia at 15 mpc, by convolving the synthetic spectra 206 expected hard x-ray and soft gamma-ray from supernovae 10 10020 50 200 1 0 − 8 1 0 − 7 1 0 − 6 1 0 − 5 1 0 − 4 n o rm a liz e d c o u n ts s k e v energy (kev) dd2d_asym_04_dc2 15mpc 10 10020 50 200 1 0 − 8 1 0 − 7 1 0 − 6 1 0 − 5 1 0 − 4 n o rm a liz e d c o u n ts s k e v energy (kev) w7 15mpc 10 10020 50 200 1 0 − 8 1 0 − 7 1 0 − 6 1 0 − 5 1 0 − 4 n o rm a liz e d c o u n ts s k e v energy (kev) dd2d_iso_04_dc3 15mpc figure 2: detector response simulations for an exposure of 106 seconds for selected models (tab. 1), for hxi (black) and sgd (red) on board astro-h. the model spectra at 20 days after the explosion are used as input, placed at distances of 15 mpc. the sensitivity curves are adopted from kokubun et al. (2010), tajima et al. (2010), and takahashi et al. (2010). note that the photon count is very low in the hxi band for all the models at this distance, thus an apparent detection by hxi (left panel) just comes from the statistical fluctuation. table 1: expected detectability (for an exposure of 106 s centered at the peak date in each band pass). shown here are limiting distance and the expected number of sne ia within the distance (shown in parenthesis). ‘cons’ and ‘opt’ are conservative and optimistic estimates, respectively. see maeda et al. (2012) for details. dd2d asym 04 w7 dd2d iso 04 m(56ni)/m� 1.02 0.64 0.42 band (kev) instrument mpc (sne year−1) 60–80 hxi 13.9 (0.43) 17.7 (0.96) 10.5 (0.09) nustar (cons.) 13.0 (0.43) 16.5 (0.70) 9.7 (0.09) nustar (opt.) 18.4 (1.13) 23.3 (2.52) 13.8 (0.43) 158 spi 4.6 (<0.09) 2.9 (<0.09) 2.3 (<0.09) sgd (cons.) 22.2 (2.09) 14.2 (0.43) 11.4 (0.09) sgd (opt.) 38.5 (6.70) 24.6 (2.96) 19.7 (1.57) 200–460 spi 3.7 (<0.09) 2.7 (<0.09) 2.3 (<0.09) sgd (cons.) 11.6 (0.09) 8.6 (0.09) 7.1 (0.09) sgd (opt.) 20.2 (1.74) 14.8 (0.43) 12.3 (0.26) 812 spi 4.3 (<0.09) 2.6 (<0.09) 2.0 (<0.09) grips 16.8 (0.87) 10.0 (0.09) 7.6 (0.09) 847 spi 7.7 (0.09) 5.4 (<0.09) 4.6 (<0.09) grips 29.8 (4.52) 21.0 (2.00) 18.0 (1.04) and designed sensitivity curves of hxi and sgd (here the adopted sensitivity curve corresponds to our ‘optimistic’ case). in 2011-2012, 6 sne ia were discovered within ∼ 20 mpc, 3 of which were within ∼ 15 mpc (from the asiago sn catalog; barbon et al., 1999). most of these were discovered soon after the explosion, and especially the nearest ones were all discovered within a week after the explosion – thus, too follow-up at the hard x and soft γ-ray peak (2 3 weeks after the explosion) is feasible. with 106 s exposure, we predict that a few (optimistic) or one (conservative) sne ia per year are reachable by astro-h. 207 keiichi maeda, yukikatsu terada, aya bamba 100 1000 1e-8 1e-7 1e-6 sgd hxi (b) 60 day fl ux [c m 2 s 1 k ev 1 ] energy (kev) 100 1000 1e-8 1e-7 1e-6 sgd hxi (a) 20 day fl ux [c m 2 s 1 k ev 1 figure 3: examples of synthetic spectra at (a) 20 days and (b) 60 days after the explosion. shown here are angleaveraged spectra for models dd2d asym 04 (red line), w7 (gray; nomoto et al., 1984), and dd2d iso 04 (dark blue) (at 10 mpc). the angle-dependent spectra seen from two opposite directions are shown for dd2d iso 04 (green and cyan) at 20 days. at 60 days the angle dependence is small. the angle dependence is small for dd2d asym 04 at both epochs. sensitivity curves for an exposure with 106 seconds of hxi and sgd on board astro-h (tajima et al., 2010; takahashi et al., 2010) are shown by black lines. 4 discussion and conclusions according to our simulations of radioactive decay signals from sne ia, the new generation hard x-ray and soft γ-ray observatories (either nustar or astro-h) are expected to be capable of detecting these signals from sne ia up to ∼ 20 mpc with 106 s exposure. this will hopefully lead to nearly annual detections, dramatically changing the field. we thus propose follow-up of nearby sne ia by these telescopes in a too mode. with a standard set up with a few 105 s exposure, the detection will be limited to extremely nearby objects (i.e., up to ∼8 12 mpc for sne ia with average brightness), but still there is a good chance of first solid detection of the signal from sne ia. once detected, it will provide various diagnostics on explosive nucleosynthesis and explosion mechanisms, and here a combination of hard x-ray and soft γ-ray will be essential. a similar argument applies for core-collapse sne. by combining results from similar simulations for corecollapse sne (maeda, 2006) and those obtained for sne ia (maeda et al., 2012), we find the following. sne iip (an explosion of a red supergiant) is not a promising target in the soft bands, since the thick h envelope is still opaque in the early phase where the 56ni decay can provide the strong emission in these band passes. among different types of core-collapse sne, sne iib/ib/ic (an explosion of a he or c+o star) are most promising. we estimate that the peak date in the high energy emission will be similar to (or a bit delayed than) that of sne ia (i.e., 2 3 weeks since the explosion). the smaller amount of 56ni here (typically ∼ 0.1m�), taking into account the delayed peak date, results in the expected horizon as 5 − 8 mpc (conservative and optimistic). such nearby objects are rare, but expected to be discovered at a rate of a few in a decade according to the past statistics. acknowledgement km thanks franco giovannelli and the organizers of frascati workshop 2013 for creating the friendly and stimulating atmosphere. the work by km has been supported by wpi initiative, mext, japan. the authors acknowledge financial support by grant-in-aid for scientific research from mext (22684012, 23340055, 23740141). references [1] barbon, r., et al.: 1999, a&as, 139, 531 [2] dotani, t., et al.: 1987, nature, 330, 230 doi:10.1038/330230a0 208 http://dx.doi.org/10.1038/330230a0 expected hard x-ray and soft gamma-ray from supernovae [3] greiner, j., et al.: 2012, exp. astron., 34, 551 doi:10.1007/s10686-011-9255-0 [4] isern, j., et al.: 2013, a&a, 552, 97 [5] kasen, d., röpke, f.k., woosley, s.e.: 2009, nature, 460, 869 doi:10.1038/nature08256 [6] koglin, j.e., et al.: 2005, proc. spie, 5900, 266 [7] kokubun, m., et al.: 2010, proc. spie, 7732, 33 doi:10.1117/12.857933 [8] khokhlov, a.: 1991, a&a, 245, 114 [9] maeda, k.: 2006, apj, 644, 385 [10] maeda, k., et al.: 2010a, nature, 466, 82 doi:10.1038/nature09122 [11] maeda, k., et al.: 2010b, apj, 712, 624 doi:10.1088/0004-637x/712/1/624 [12] maeda, k., et al.: 2012, apj, 760, 54 doi:10.1088/0004-637x/760/1/54 [13] milne, p.a., et al.: 2004, apj, 613, 1101 doi:10.1086/423235 [14] nomoto, k., thielemann, f.-k., yokoi, k.: 1984, apj, 286, 644 [15] roques, j.p., et al.: 2003, a&a, 411, l91 [16] seitenzahl, i.r., et al.: 2013, mnras, 429, 1156 [17] summa, a., et al.: 2013, a&a, 554, 67 [18] sunyaev, r.a., et al.: 1987, nature, 330, 227 doi:10.1038/330227a0 [19] tajima, t., et al.: 2010, proc. spie, 7732, 34 [20] takahashi, t., et al.: 2010, proc. spie., 7732, 27 209 http://dx.doi.org/10.1007/s10686-011-9255-0 http://dx.doi.org/10.1038/nature08256 http://dx.doi.org/10.1117/12.857933 http://dx.doi.org/10.1038/nature09122 http://dx.doi.org/10.1088/0004-637x/712/1/624 http://dx.doi.org/10.1088/0004-637x/760/1/54 http://dx.doi.org/10.1086/423235 http://dx.doi.org/10.1038/330227a0 introduction expected characteristics observational perspectives discussion and conclusions 156 acta polytechnica ctu proceedings 2(1): 156–160, 2015 156 doi: 10.14311/app.2015.02.0156 outburst properties of v1504 cyg and v344 lyr j. k. cannizzo1,2 1cresst, dept. of phys., umbc, baltimore, md 21250, usa 2nasa/gsfc/astroparticle physics lab., greenbelt, md 20771, usa corresponding author: john.k.cannizzo@nasa.gov abstract i begin by reviewing dwarf novae and the disk instability theory, and then present an overview of three ideas for producing superoutbursts in the su uma stars − the thermal tidal instability, irradiation-induced secondary mass overflow, and the plain vanilla disk limit cycle instability. i discuss the properties of the outbursts in two su uma systems observed by kepler in the context of the three theories. i conclude with a look beyond the su uma systems. keywords: cataclysmic variables dwarf novae optical photometry individual: v1504 cyg, v344 lyr, u gem, ss cyg. 1 introduction 1.1 disk instability model the outbursts in dwarf novae (dne), a subclass of cataclysmic variables (cvs warner 1995), are thought to be due to a limit cycle instability operating within the accretion disk (smak 1984) that episodically stores up matter (during quiescence) and dumps a fraction of it onto the accreting wd (during outburst). the basic disk instability model is now generally accepted as the correct explanation for dwarf nova outbursts (smak 1984, cannizzo 1993a, lasota 2001). the strongest point of the model is its prediction of a dividing line between novalike and dwarf nova systems (smak 1984), a line which has no free parameters1. for the systems above this line, the rate of accretion feeding into the outer disk is such that the entire disk is always in a hot, ionized state. other strengths of the theory include its natural ability to account for slow and fast decays of dn outbursts as being due either to viscous or thermal decay, the natural tendency of the model to produce alternating long and short outbursts (cannizzo 1993b), and the fact that the outer disk radius router continues to contract in quiescence following a dn outburst, whereas in the secondary mass-overflow model there is a one-to-one relation between disk luminosity and router (ichikawa & osaki 1992). smak (1984) was the first to use the relation between the fast rate of decay in dwarf nova outbursts and orbital period (i.e., the bailey relation) to constrain the value of the shakura-sunyaev (1973) alpha parameter in the ionized disk, αhot ' 0.1. the high fidelity kepler data now enable us to trace deviations from strictly exponential decay for the fast decays (cannizzo et al. 2012=c12), and also to utilize the slow rate of decay, wherein the entire disk resides in a hot, ionized state. the fact that the viscous and thermal time scales differ by a factor h/r means that this ratio can be directly calculated to be ∼0.01 − 0.02 (cannizzo 1998; menou, hameury, & stehle 1999). although successful as a model for outbursts, it has been known for some time that there must be various add-ons to the theory in order to have a complete theory for the accretion disk, including something like a coronal siphon flow (meyer & meyer-hofmeister 1994) or some other type of evaporative instability (e.g., shaviv & wehrse 1986) to transfer matter from the disk onto the wd during quiescence to get accretion. 1.2 su uma systems the su uma stars represent short orbital period dwarf novae (dne), porb < 2 hr, displaying normal outbursts (nos) and superoutbursts (sos) long outbursts with superhumps (shs). although the contrast between nos and sos seems quite dramatic, van paradijs (1983) has shown that the properties of sos, seen in systems below the period gap, can be understood as a natural continuation of the long outbursts in dne above the period gap porb > 3 hr. in addition, sos are defined by the existence of shs, modulations in the outburst 1 for ∼10 yr it appeared that the best studied dwarf nova, ss cyg, was on the wrong side of the line (harrison et al. 1999, schreiber, hameury, & lasota 2003). however, this prognosis was based on an incorrect distance (159 ± 12 pc). the corrected distance (114 ± 2 pc) reconciles data and theory (miller-jones et al. 2013). 156 http://dx.doi.org/10.14311/app.2015.02.0156 outburst properties of v1504 cyg and v344 lyr light curve at periods a few percent longer than the orbital period. the prevailing theory is that when the outer disk edge surpasses the point of 3:1 resonance with the binary orbit, an eccentric disk develops that precesses progradely in the binary frame (whitehurst 1988). more recent work indicates that the disk is not statically eccentric but has a fundamental “breathing” mode − oscillating between circular and eccentric (montgomery 2012ab). these shs are referred to as positive superhumps (pshs) since their periods exceed porb; there also exist negative superhumps (nshs) with periods a little less than porb thought to be indicative of a tilted, retrogradely precessing disk (patterson et al. 1993). the two disk states can and sometimes do coexist (montgomery 2012ab). in recent years three scenarios have emerged to account for sos: (i) the thermal-tidal instability (tti osaki 1989) which combines the normal disk instability model and the precessing disk model, (ii) irradiationinduced enhanced mass overflow from the mass-losing star, and (iii) the plain disk instability (pdi) model (cannizzo et al. 2010=c10). in a recent series of papers osaki & kato (2013abc) have discussed the relative merits of these three models within the context of recent kepler observations of v1504 cyg and v344 lyr. 2 kepler observations of v1504 cyg and v344 lyr c12 find that the durations of the nos within a supercycle increase monotonically, whereas the durations of the quiescence intervals between nos can either increase monotonically, or else increase to a local maximum about half way to the next so, and then decrease. these supercycles for which the recurrence times for nos reach a maximum roughly half way between sos present a challenge for the tti model, in which one expects a monotonic increase in disk mass, angular momentum, and triggering radius as the next so approaches. osaki & kato (2013c) acknowledge this difficulty, and also point out that this criticism, given by c12, applies equally well to the pdi calculations presented by c10. this is a good point − in fact the calculations presented in c10 did not show any variation in recurrence time for nos. figure 1 shows the latest pdi calculations of sos in a su uma. ok13a discuss the amplitude of the precursor dip and note that, at shorter wavelengths, the dip tends to stand out more strongly. one of the best examples is examples is in oy car (porb = 1.51 hr; mauche & raymond 2000) where simultaneous visual and euv (∼100 å) observations are presented of a so. embedded precursors which are simultaneously small in the optical and large in the euv provide a challenge to both tti and pdi models. in the tti model, the amplitude of the precursor can be precisely controlled by tweaking the strength of the increase in tidal torque accompanying the crossing of the 3:1 resonance radius by router, and the delay time ttti, delay accompanying its onset. for the sos there can be no precursors, as in the early tti models (e.g., ichikawa, hirose, & osaki 1993), or there can be precursors which are well-defined both in v and euv wavelengths, e.g., in models with ttti, delay = 1.75 d (schreiber, et al. 2004, see their fig. 3). v344 lyr figure 1: a comparison of kepler data on v344 cyg with two sos in a long light curve using the pdi model. a new idea introduced in ok13a is that the presence of time-varying nshs, when they are present, gives us a handle on variations in the outer disk radius. larwood (1998) gives a general expression for the (nodal) precessional frequency associated with a tilted disk ωnpr = ωorb −ωnsh which is based on the surface density distribution σ(r) and accretion disk tilt angle θ, ωnpr = 2πνnpr = − 3 4 gm2 a3 ∫ σr3dr∫ σωr3dr cos θ, (1) where m2 is the secondary mass and a the binary separation. for a nominal power law distribution in surface density, νnpr νorb = −η 3 7 q √ 1 + q (router a )3/2 cos θ, (2) where η is a constant of order unity. the observed νnsh(t) variations in v1504 cyg reveal a disk contraction between nos, superposed on a longer term expansion between sos. this approach was criticized by smak (2013) who argued that, if one takes the time derivative of larwood’s pnsh equation, one obtains an expression with five terms (see his eqn. [7]), only one of which is related to the outer disk radius. 157 j. k. cannizzo the disk instability model provides a definitive prediction of σ(r) at all times, and therefore one may carry out a calculation of the long-term accretion disk evolution relevant for an su uma system, and do a direction comparison of νnsh calculated using the exact and approximate expressions. the results are shown in figure 2. the panels show the light curve, disk mass, disk radius, and the exact and approximate calculations for νnsh. the approximate expression gives a very good representation of the exact formula, which vindicates the underlying idea of ok13a and supports their use of νnsh(t) to track router(t). figure 2: a calculation of the pdi model for a su uma system. shown are (i) the kepler light curve, (ii) the disk mass, in units of 1023 g, (iii) the outer disk radius, in units of 1010 cm, (iv) the nsh frequency calculated using the larwood formula, and (v) the nsh frequency calculated using the approximate formula given ok13abc, taking η = 1.5. there is actually only one small difference between osaki’s tti and the pdi model for superoutbursts, namely, in the tti model there is a strong and brief contraction of the outer disk radius due to the (hypothesized) strong increase in tidal torque from the secondary star. an obvious way to distinguish between the two models is to closely monitor the outer disk radius in an eclipsing su uma star during the early onset of a so to see if there is a sudden, brief decrease in router. smak (2013) argues that this observation has already been carried out and shows no such contraction; ok13c counter that the time period in question was well into the so and therefore missed the start of the so. 3 looking beyond the su uma systems ok13abc argue that the precursor in sos is due to the onset of the tti, whereas i have argued it is a natural consequence of the pdi. an obvious next step would be to look in detail at long dn outbursts in systems where we do not expect the tti to operate − systems above the period gap where the outer disk radius never reaches the 3:1 radius. even if one accepts the fundamental tenet of the tti, namely a causal relation between whitehurst’s precessing eccentric disk and the di, it could not possibly be a factor because the disks in long period systems are never subject to the 3:1 instability. if one sees precursors in long dn outbursts in these systems, then they must be due to a more general aspect of the di. the two best studied dne are both longward of the period gap and should not experience the 3:1 instability − u gem and ss cyg. they have > 100 yr light curves, but unfortunately the bulk of the data is not precise enough to enable any statement to be made about the detailed shape of the outburst onset. cannizzo (2012) examined the aavso data for these systems, and found brief stretches of the light curves for which the number of observations entering into daily means exceeds ∼103, concomitant with the use of digital photometry. although still not on a par with kepler data, these data do afford a more detailed look at the outburst light curve shapes than the usual data. there are two long outbursts in u gem and one in ss cyg for which a statement can be made concerning the shape of the outburst light curve during outburst onset; one of the u gem outbursts shows no obvious precursor, one does, and the one ss cyg long outburst also does. in fact, the flattened shape of the precursor is similar to that seen in fig. 13 (panel 2) of cannizzo (1993b) in modeling ss cyg. these outbursts provide a hint that embedded precursors during long outbursts may not be confined to the su uma systems. 4 summary kepler light curves of short period dne have resparked interest in the nature of sos and led to the question: is the thermal-tidal instability needed, or can the plain vanilla version of the accretion disk limit cycle do the job all by itself? a detailed time-resolved study of an eclipsing su uma system during so onset should settle the question − if there is a dramatic contraction of the disk at so onset, the tti would be preferred; if not, the plain di model would be sufficient. also, recent work 158 outburst properties of v1504 cyg and v344 lyr by osaki & kato has shown that the time varying nsh frequencies νnsh(t) can be taken as a surrogate for the outer disk radius variations router(t). finally, it may be necessary to look beyond the short period dwarf novae to gain perspective on the nature of embedded precursors in long outbursts. references [1] cannizzo, j.k.: 1993a, in accretion disks in compact stellar systems ed., j.c. wheeler (singapore: world scientific), 6. [2] cannizzo, j.k.: 1993b, apj, 419, 318. doi:10.1086/173486 [3] cannizzo, j.k.: 1998, apj, 494, 366. doi:10.1086/305210 [4] cannizzo, j.k.: 2012, apj, 757, 174. doi:10.1088/0004-637x/757/2/174 [5] cannizzo, j.k., smale, a.p., wood, m.a., et al.: 2012, apj, 747, 117. doi:10.1088/0004-637x/747/2/117 [6] cannizzo, j.k., still, m.d., howell, et al.: 2010, apj, 725, 1393. doi:10.1088/0004-637x/725/2/1393 [7] harrison, t.e., mcnamara, b.j., szkody, p., et al.: 1999, apj, 515, l93. doi:10.1086/311982 [8] ichikawa, s., hirose, h., osaki, y.: 1993, pasj, 45, 243. [9] ichikawa, s., osaki, y.: 1992, pasj, 44, 15. [10] larwood, j.: 1998, mnras, 299, l32. doi:10.1046/j.1365-8711.1998.01978.x [11] lasota, j.-p.: 2001, new astron. rev., 45, 449. doi:10.1016/s1387-6473(01)00112-9 [12] mauche, c.w., raymond, j.c.: 2000, apj, 541, 924. doi:10.1086/309489 [13] menou, k., hameury, j.-m., stehle, r.: 1999, mnras, 305, 79. [14] meyer, f., meyer-hofmeister, e.: 1994, a&a, 288, 175. [15] miller-jones, j.c.a., sivakoff, g.r., knigge, c., et al.: 2013, science, 340, 950. doi:10.1126/science.1237145 [16] montgomery, m.m.: 2012a, apj, 745, l25. doi:10.1088/2041-8205/745/2/l25 [17] montgomery, m.m.: 2012b, apj, 753, l27. doi:10.1088/2041-8205/753/2/l27 [18] osaki, y.: 1989, pasj, 41, 1005. [19] osaki, y., kato, t.: 2013a, pasj, 65, 50. [20] osaki, y., kato, t.: 2013b, astro-ph/1305.5877. [21] osaki, y., kato, t.: 2013c, astro-ph/1309.3722. [22] patterson, j., thomas, g., skillman, d.r., diaz, m.: 1993, apjs, 86, 235. [23] schreiber, m.r., hameury, j.-m., lasota, j.-p.: 2003, a&a, 410, 239. [24] schreiber, m.r., hameury, j.-m., lasota, j.-p.: 2004, a&a, 427, 621. [25] shakura, n.i., sunyaev, r.a.: 1973, a&a, 24, 337. [26] shaviv, g., wehrse, r.: 1986, a&a, 159, l5. [27] smak, j.: 1984, acta astron., 34, 161. [28] smak, j.: 2013, acta astron., 63, 109. [29] van paradijs, j.: 1983, a&a, 125, l16. [30] warner, b.: 1995, cataclysmic variables (cambridge: cambridge univ. press). [31] whitehurst, r.: 1988, mnras, 232, 35. discussion margaretha pretorius: if there is no tidal instability how do you make superhumps? john cannizzo: the tidal instability is not being disputed. in both osaki’s thermal-tidal instability model and the plain disk instability model superhumps are due to a precessing eccentric disk which becomes manifest when the outer disk radius surpasses the the 3:1 radius. the question is then whether this has anything to do with the superoutbursts: osaki would say yes, i would say no. raymundo baptista: yesterday i showed inferred αcold values for three quiescent su uma stars, which show αcold ∼ 10−1 ∼ αhot. my interpretation is that this eliminates the possibility of the disc instability model to explain their outbursts. could you please comment on this? john cannizzo: what’s the temperature of that gas? raymundo baptista: about 6000 k. 159 http://dx.doi.org/10.1086/173486 http://dx.doi.org/10.1086/305210 http://dx.doi.org/10.1088/0004-637x/757/2/174 http://dx.doi.org/10.1088/0004-637x/747/2/117 http://dx.doi.org/10.1088/0004-637x/725/2/1393 http://dx.doi.org/10.1086/311982 http://dx.doi.org/10.1046/j.1365-8711.1998.01978.x http://dx.doi.org/10.1016/s1387-6473(01)00112-9 http://dx.doi.org/10.1086/309489 http://dx.doi.org/10.1126/science.1237145 http://dx.doi.org/10.1088/2041-8205/745/2/l25 http://dx.doi.org/10.1088/2041-8205/753/2/l27 j. k. cannizzo john cannizzo: one of the limitations of the standard disk instability theory has always been that it’s not a complete theory certainly it doesn’t work at all to describe the quiescent state. if you take the model seriously, the plots in the movie i showed indicate temperatures in quiescence of ∼2000-3000 k. also, we see significant evidence for accretion during quiescence mainly through soft x-ray emission, which is not predicted by the di theory since in quiescence the accretion rate would be zero. so the bottom line is that the quiescence gas you observe must be due to something extrinsic to the di model. 160 introduction disk instability model su uma systems kepler observations of v1504 cyg and v344 lyr looking beyond the su uma systems summary 183 acta polytechnica ctu proceedings 2(1): 183–187, 2015 183 doi: 10.14311/app.2015.02.0183 close-in substellar companions and the formation of sdb-type close binary stars l. y. zhu1,2,3, s. b. qian1,2,3, e.-g. zhao1,2, e. fernández lajús4,5, z.-t. han1,2,3 1yunnan observatories, chinese academy of sciences, p.o. box 110, 650011 kunming, p.r. china 2key laboratory for the structure and evolution of celestial bodies, chinese academy of sciences, p.o. box 110, 650011 kunming, p. r. china. 3physical sciences department, graduate university of the chinese academy of sciences, yuquan road 19, sijingshang block, 100049 beijing, china. 4facultad de ciencias astronómicasy geof́ısicas, universidad nacional de la plata, 1900 la plata, buenos aires, argentina. 5instituto de astrofisica de la plata (cct la plata conicet/unlp), argentina. corresponding author: zhuly@ynao.ac.cn abstract the sdb-type close binaries are believed to have experienced a common-envelope phase and may evolve into cataclysmic binaries (cvs). about 10 % of all known sdb binaries are eclipsing binaries consisting of very hot subdwarf primaries and low-mass companions with short orbital periods. the eclipse profiles of these systems are very narrow and deep, which benefits the determination of high precise eclipsing times and makes the detection of small and close-in tertiary bodies possible. since 2006 we have monitored some sdb-type eclipsing binaries to search for the close-in substellar companions by analyzing the light travel time effect. here some progresses of the program are reviewed and the formation of sdb-type binary is discussed. keywords: hot subdwarf binary photometry individual: ny vir, hs 0705+6700, nsvs 07826147. 1 introduction hot subdwarf stars (sdb) lie in the extreme horizontal branch (ehb). they are burning helium in their cores and have very thin hydrogen envelopes. after leaving the main sequence, the progenitors of sdbs would evolve to red giant, ignite helium and settle down on the ehb (heber 2009). they have high temperatures and gravities. to explain these characters, their progenitors must have experienced enhanced mass loss at the tip of red giant branch (castellani and castellani, 1993; d’cruz et al., 1996). however, what causes this extreme mass loss remains an open question. many authors suggested that the mass loss is triggered by the presence of a companion. in 2007, silvotti et al., found a giant planet orbiting around a sdb star, v391 pegasi, at a distance of 1.7 au. after that, an even more extreme case was published by charpinet et al., (2011). they reported the presence of two nearly earth-sized bodies orbiting the sdb star kic 05807616 at distances of 0.0060 au and 0.0076 au, with orbital periods of 5.7625 and 8.2293 hours, respectively. these detections support the way to form single sdb stars that these bodies were originally giant planets immersed in the red-giant envelope and massive enough to survive engulfment and triggered the enhanced mass loss necessary for the formation of an sdb star. so the close-in substellar companions paly an important role to form single sdb stars. actually, about half (or may be more) of the sdbs reside in close binaries (morales-rueda et al., 2006). are there any substellar objects around sdb close binaries just like the substellars founded around the single sdb stars? recently, some circumbinary substellars orbiting around sdb eclipsing binaries were indeed detected, such as hw vir (lee et al., 2009), hs 0705+6700 (qian et al, 2009), ny vir (qian et al., 2012) and sdssj0820+0008 (geier et al, 2012). the sdb close binaries are believed to be formed from binary systems through common envelope(ce) ejection (han et al., 2002, 2003). because the separation of the two components in sdb close binaries is much less than the size of the subdwarf progenitor in its red-giant phase, these systems must have experienced a ce phase. after spiral in and envelope ejection, the present binary system consisting of a sdb star and a low-mass companion star was formed with short periods. the presence of substellar tertiary, especially the close-in objects, orbiting sdb close binaries provides some clues to the inter183 http://dx.doi.org/10.14311/app.2015.02.0183 l. y. zhu et al. active effect between them, and may give evidences of the formation of the planets and the sdb close binaries. 2 targets and method for searching for close-in substellar companions of the sdb close binaries, we chose eclipsing sdb close binaries (i.e. hw vir binaries) as our targets to monitor because their eclipsing profiles provide us with very useful information about their orbital period variations, which may be caused by the presence of third bodies. hw vir binaries have been known for more than four decades. they are detached eclipsing binaries composed of a sdb star and a low-mass companion star. till now, not more than fifteen of such objects were discovered. their well-detached configuration indicates that their eclipses might not be influenced by the accretion-driven radiations from other parts (e.g., the accretion discs and hot spots etc.) observed in cvs. and the compact structures of both components make the eclipses in the light curves very narrow and deep. these characteristics favour a highly precise determination of their times of light minimum, generally with errors less than 0.0001days. thus very small amplitude orbital period variations due to the substellar tertiaries can be detected by analyzing the observed-calculated (o-c) diagrams. therefore they are the most hopeful targets for detecting substellar objects around them with light time effect. using 2.4-m, 1-m and 60-cm telescopes in yunnan observatories, 2.16-m and 85-cm telescopes in xinglong station of national observatories and 2.15-m jorge sahade telescope in argentina, we began to monitor hw vir binaries since 2006. our targets are listed in table 1. table 1: summary of our targets name coordinate period mag. (j2000) h (mag) hw vir 12 44 20.24 -08 40 16.8 2.801 v=11.2 ny vir 13 38 48.14 -02 01 49.2 2.426 v=14.2 hs 0705+6700 07 10 42.09 +66 55 44.0 2.296 b=14.1 hs 2231+2441 22 34 20.89 +24 57 00.4 2.654 v=14.1 nsvs 14256825 20 20 00.51 +04 37 55.6 2.649 v=14.3 bulsc 16-335 14 45 20.21 -06 44 03.2 3.001 v=16.3 nsvs 07826147 15 33 49.40 +37 59 28.6 3.883 v=13.6 sdssj 0820+0008 08 20 53.5 +00 08 53.4 2.304 g=14.9 2m 1938+4603 19 38 32.61 +46 03 59.1 3.024 g=11.9 ec 10246-2707 10 26 56.47 -27 22 57.1 2.844 b=14.2 asas 102322-3737 10 23 21.89 -37 36 59.9 3.343 v=11.6 supposing that there is a third body existing in this system, under the mutual gravitation force, the components of eclipsing pair would rotate around the barycenter of this triple system. seen by a distant observer, the light-travel time of this system will change because of the change in the orbital movement caused by the additional body, which can result in the observed cyclic change in the o-c diagram constructed by the eclipsing timings. this light time effect is very useful to detect tertiaries within au distance level in evolved eclipsing binaries compared with the radial velocity and transit methods, which have been extensively used to search for planets around solar-type main-sequence stars. that is because the high surface gravity of the evolved stars prevents us to achieve the high radial velocity precision required to detect planets, and the small size of the components (for example, r1 = 0.166 and r2 = 0.152 for nsvs 07826147 derived by for et al., 2010.) make the low probability of transits, see silvotti (2009) for more information. 3 new results here we present new results of our three targets (hs 0705+6700, ny vir and nsvs 07826147) from our ongoing search for circumbinary substellars orbiting around sdb eclipsing binaries. for hs 0705+6700 and ny vir, new observations confirmed our previous study (qian et al., 2009, 2012). for nsvs 07826147, preliminary analysis of the o-c diagram constructed by our five years data implied the presence of a close-in planet orbiting around it at a distance about 0.64au. 3.1 hs 0705+6700 hs0705+6700 ( = gsc 4123-265) was listed as a dwarf candidate from the hamburg schmidt survey (hagen et al. 1995 ). the observations indicated that it is an eclipsing binary. a detailed photometric and spectroscopic investigation was carried out by drechsel et al. (2001). absolute parameters of both components were determined suggesting that the primary is an sdb star, while the secondary is a cool stellar object. a cyclic change in the o-c curve of hs 0705+6700 was discovered by qian et al. (2009), which indicated the presence of a possible brown dwarf companion. it was later confirmed by camurdan et al. (2012), beuermann et al. (2012), and qian et al. (2013) who revised the parameters of the brown dwarf. the (o − c)1 diagram of the sdb-type binary hs 0705+6700 with respect to the following linear ephemeris, min.i = bjd 2451822.760549 + 0.09564665 × e, (1) was displayed in fig. 1, where dots refer to our new data obtained after qian et al. (2009). it is shown that, apart from the cyclic change reported by previous authors, there is an upward parabolic variation in the (o −c)1 curve. this reveals a period increase at a rate of ṗ = +9.8×10−9 days/year. the angular momentum loss via gravitational radiation or/and magnetic braking should cause a decrease in the orbital period rather than increase. moreover, the period increase cannot be 184 close-in substellar companions and the formation of sdb-type close binary stars 0 10000 20000 30000 40000 50000 -0.003 -0.002 -0.001 0.000 0.001 0.002 0.003 e figure 1: o-c diagram of hs 0705+6700. dots refer to our new data obtained after qian et al. (2009), and open circles to the other data. dashed line refers to the continuous increase in the period, while the solid line to a combination of the period increase and a cyclic change. explained by mass transfer between the components because of the detached configuration of the system. therefore, the observed upward parabolic variation may be only a part of another cyclic change with very longperiod that is caused by the light-travel time effect due to the presence of another substellar object in a wider orbit. 3.2 ny vir ny vir (pg 1336-018) was discovered as a hw vir binary by kilkenny et al., (1998) which is composed of 2.580 2.585 2.590 2.595 2.600 2.605 0.4 0.2 0.0 -0.2 -0.4 6.750 6.755 6.760 6.765 0.4 0.2 0.0 -0.2 -0.4 figure 2: eclipse profiles of ny vir observed in november 2012 and april 2013 with the 2.15-m ”jorge sahade” telescope. dots represent the magnitude differences between the ny vir and the comparison star. a rapid pulsator and an m5-type star. kilkenny et al., (2011) found the orbital period of the binary is decreasing at a rate of ṗ = −11.2 × 10−13 d per orbit. a cyclic change was discovered to be superimposed on the long-term decrease by qian et al. (2012), which was explained by the presence of a circumbinary planet. qian et al. (2012) pointed out that the long-term period decrease can not be explained by angular momentum loss via gravitational radiation or/and magnetic braking and proposed that there is another substellar object in a wider orbit as in the case of hs 0705+6700. new observations of the sdb binary were obtained. the eclipse profiles obtained by using the 2.15-m ”jorge sahade” telescope in argentina were shown in fig. 2. the new determined eclipse times confirm the presence of the circumbinary planet and the parameters will be revised by including more times of light minimum. 3.3 nsvs 07826147 nsvs07826147 (2m 1533+3759) has been suspected as a possible eclipsing sdb binary system by kelley and shaw (2007) using the data from the northern sky variability survey and the two micron all sky survey. they derived the period of this system as 0.16177 days. zhu and qian (2009) improved its period as 0.16177046 days based on all available times of light minimum until that time. the detailed absolute parameters derived by the combination of photometric and spectroscopic observations were published by for et al., (2010). till now, we have monitored this target for more than 5 years. new light curves obtained with the 85-cm telescope and the 2.4m telescope are displayed in fig. 3. the constructed o-c diagram shows the trend of the periodic variation, which implies the presence of a closein circumbinary planet in nsvs 07826147 at a distance about 0.64au. the detailed analysis is in progress. 72.34 72.35 72.36 72.37 0.9 0.0 -0.9 m hjd2455900+ 74.256 74.260 74.264 74.268 74.272 74.276 2 1 0 -1 figure 3: light curves of nsvs07826147. left panel: eclipsing profiles of nsvs07826147 obtained with 85cm telescope on may 26, 2012 in r band. right panel: obtained with 2.4-m telescope on feb. 14, 2012 without filters. dots refer to the magnitude differences between the target and the comparison star, and open circles to the magnitude differences between the comparison and check star. 185 l. y. zhu et al. 4 summary one formation method of sdb-type binaries in theory is through a common envelope (ce) phase after the more massive star in the initially detached system evolves into a red giant. the ejection of the ce results in a large amount of angular momentum loss, and then a shortperiod sdb-type binary is formed. in previous section, new results of the three sdb-type eclipsing binaries are introduced. the presence of substellar companions especially the one orbiting nsvs 07826147 with the closest distance will give some constrains on the formation of this type of objects as well as on the interaction between red giants and their companions. to improve the results, more observations are required in the future. acknowledgement this work is supported by chinese natural science foundation (no.11133007 and 11325315) and yunnan natural science foundation (no. 2013fb084). new ccd photometric observations were obtained with the 2.16m and 85-cm telescope at xinglong station, the 2.4m, 1.0-m, and 60-cm telescopes at yunnan observatories in china, and the 2.15-m ”jorge sahade” telescope in argentina. references [1] beuermann, k., breitenstein, p., debski, b., diese, j., dubovsky, p. a., dreizler, s., et al., 2012, a&a 540, 8 [2] castellani, m., castellani, v., 1993, apj, 407, 649 [3] camurdan, c. m., zengin camurdan, d., ibanoǧlu, c., 2012, newa 17, 325 doi:10.1016/j.newast.2011.08.004 [4] charpinet, s., fontaine, g., brassard, p., et al., 2011, natur, 480, 496 doi:10.1038/nature10631 [5] d’cruz, n. l., dorman, b., rood, robert, t., o’connell, r. w., 1996, apj, 466, 359 doi:10.1086/177515 [6] drechsel, h., heber, u., napiwotzki, r., ostensen, r., solheim, j.-e., johannessen, f., schuh, s. l., deetjen, j., zola, s., 2001, a&a 379, 893 [7] for, b. q., et al., 2010, apj, 708, 253 doi:10.1088/0004-637x/708/1/253 [8] hagen, h-j., groote, d., engels, d., & reimers, d. 1995, a&as, 111, 195 [9] han, z. et al., 2002, mnras, 336, 449 [10] han, z. et al., 2003, mnras, 341, 669 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et al., 2007, natur, 449, 189 [23] zhu, l. y., qian, s. b., 2010, ap&ss, 329, 107 discussion raymundo baptista: light time travel effect is not the only interpretation for physical period changes in binary stars. an alternative is the magnetic activity cycle of the late-type star modulating the binary period. it is hard to believe a planet at 1 au orbit would survive commom envelope phase of a binary which was at comparable separation before. liying zhu: firstly, the cyclic period change caused by light time effect is strictly periodic change in o-c diagram not only quasi-periodic variation aroused 186 http://dx.doi.org/10.1016/j.newast.2011.08.004 http://dx.doi.org/10.1038/nature10631 http://dx.doi.org/10.1086/177515 http://dx.doi.org/10.1088/0004-637x/708/1/253 http://dx.doi.org/10.1046/j.1365-8711.2003.06451.x http://dx.doi.org/10.1146/annurev-astro-082708-101836 http://dx.doi.org/10.1111/j.1365-2966.2010.17919.x http://dx.doi.org/10.1046/j.1365-8711.1998.01432.x close-in substellar companions and the formation of sdb-type close binary stars by the magnetic activity cycle. secondly, in our fiveyears monitor, no asymmetric light curves caused by the magnetic activity for these hw vir binaries were found. their light curves are very stable. thirdly, one of our target sdss j0820+0008 is composed of an sdb star and a brown dwarf, the cyclic change can not be explained by magnetic activity cycles. the detected planet at 1 au orbit may be protected by the engulfment of the low mass stellar companion during the ce phase. and the other possibility is that the planet is the second-generation one. 187 introduction targets and method new results hs0705+6700 ny vir nsvs07826147 summary 133 acta polytechnica ctu proceedings 2(1): 133–138, 2015 133 doi: 10.14311/app.2015.02.0133 observing cvs and lmxbs with salt: updates and recent results david a. h. buckley1 1southern african large telescope, c/o south african astronomical observatory, po box 9, observatory 7935, cape town, south africa corresponding author: dibnob@salt.ac.za abstract i present an overview of the ongoing observational programs utilizing the southern african large telescope (salt), focussing on magnetic cataclysmic variables and low mass x-ray binaries. salt’s instruments and capabilities are well suited to time resolved studies of the accretion phenomena exhibited in these systems. initial observations, using salticam, have been used to derive high time resolution (∼100ms) eclipse light curves, with high signal-to-noise, of polars. recently this work has been extended to time resolved spectroscopic studies, utilizing the salt robert stobie spectrograph (rss), allowing an opportunity to probe how the emission lines change during eclipse. a program to search for and characterize quasi-periodic oscillations (qpos) in magnetic cvs using the saao 1.9-m telescope, begun in 2012, has been expanded to include observations with a photon counting camera (bvit) on salt. a multi-longitude campaign involving salt, other saao facilities plus the eso ntt, was carried out in march/april 2012 on the enigmatic gamma ray source, xss j12270–4859, which has revealed the that system exhibits spectral line variations, from absorption to emission, seemingly over timescales of < 1 h. keywords: cataclysmic variables polars intermediate polars optical spectroscopy photometry x-rays individual: fl cet, ex hya, hu aqr, xss j12270-4859. 1 introduction the southern african large telescope (salt) is one of five 10-m class segmented mirror telescopes and the only one situated in the southern hemisphere. at the last palermo cv meeting in 2011, i summarized the capabilities and the status at that time (buckley 2012) and presented some early science results. since then salt has entered full time science operations (in november 2011) and has begun to steadily increase its scientific productivity (up to ∼60 refereed publications). there are still some remaining issues to resolve, however, notably implementation of a mirror edge sensing system to allow active control of the primary mirror array. this is currently under development and the first batch of new inductive edge sensors will be installed early in 2014, while the final ones will be installed late in 2015. this will lead to much improved image quality and, more importantly its retention during an observation. similarly there are improvements being made to the throughput of the main work-horse salt science instrument, the robert stobie spectrograph (rss), which should greatly improve its performance, particularly in the blue (320 400nm). this is being achieved by fabrication of new collimator optics and opto-mechanics, designed to avoid the optical fluid contamination problems that have led to the current throughput under-performance. these new optics are expected to be installed by mid-2014. salt is a 100% queue-scheduled telescope, where all observations are undertaken in a service observing mode by salt staff. this modus operandi lends itself to target-of-opportunity observations and synoptic monitoring campaigns, both of which are often appropriate for observations of cataclysmic variables and related objects. for example, to cover specific phase intervals, to coordinate with x-ray or other satellite observations, to monitor a system through an outburst or accretion state change, etc. just such observations have already been carried out with salt and some examples are discussed in this paper. 1.1 high speed photometry with salticam salticam employs frame transfer (ft) ccds, which allows for fast acquisition and imaging. it also has a moving occulting mask which allows for 3 imaging modes: full field imaging, frame transfer mode (half of the field is imaged) and a specialized high-speed mode, 133 http://dx.doi.org/10.14311/app.2015.02.0133 david a. h. buckley where a narrow slot is imaged. the latter, known as ”slot mode”, allows for high-speed (up to ∼12.5hz) photometry and has been used extensively for observing cvs, x-ray binaries and related systems. figure 1: the 11-m southern african large telescope (salt), consisting of 91 × 1.2 m (corner-to-corner) hexagonal mirror segments. this mode allows for just 144 rows of the ccd mosaic to be illuminated, equivalent to ∼20 arcsec × 10 arcmin rectangular region on the sky. the object of interest, plus ideally a nearby (< a few arcmin) constant source (i.e. a comparison star), are placed in the centre of this slot, which can be achieved by rotating the instrument to the desired position angle. the illuminated region of the ccd is then exposed (typically with exposure times ranging from 80 – 1000 ms) and then quickly (within a few ms) row-shifted into the storage array, while another exposure begins. thus exposed slot regions are sequentially moved through the storage array to the serial readout register. during the next exposure, the latest exposed region to reach the bottom of the chip is readout in a time < the exposure time, which is dependent on the on-chip binning factor (usually 4 × 4 pixels). 1.2 time resolved spectroscopy with rss like for all current salt instruments, the robert stobie spectrograph (rss) resides at salt’s prime focus (f/4.2). the advantages of this include maximizing the performance at short wavelengths (< 400 nm), by avoiding use of optical fibres, and ability to conduct fabryperot imaging spectroscopy. rss was designed to have a range of capabilities and observing modes, each one remotely and rapidly reconfigurable. figure 2: the narrow occulting slot that can be moved just in front of the salticam ccds to allow for high time resolution (down to ∼80 ms) photometry. in keeping with the overall philosophy of exploiting those areas where salt has a competitive edge, the instrument has a range of capabilities, many well suited to cv and lmxb studies, including: � wavelength coverage from 320 to 900 nm for the initial uv-vis configuration (a second near ir arm, extending to 1.7 µm, is under construction). low to medium resolution spectroscopy (up to r ∼6000 with 1 arcsec slits) using efficient and tuneable vph gratings. � narrow band, tuneable filter and fabry-perot imaging. � high time resolution (∼0.1 sec) spectroscopic and polarimetric modes. � imaging and spectropolarimetry (linear, circular and all stokes modes), time resolved modes plus low resolution (r ∼50) wide-field imaging spectropolarimetry. high time resolution spectroscopy is achieved using a similar slot-mode option, as for salticam. however, rather than the mask being just above the detector, it is placed at the telescope’s focal plane, coplanar with the spectrograph’s slit. this allows for the spectrum of the variable target to be imaged in a narrow rectangular region and then shuffled into the storage array at the end of the exposure. three 2k × 4k ccds are mosaiced to allow spectra to be obtained over the entire 8 arcmin fov (e.g. for mos). figure 3 shows the arrangement of 134 observing cvs and lmxbs with salt:updates and recent results the slot with respect to the ccd mosaic, where the slot is placed just above the imaging / storage boundary. figure 3: schematic showing the region on the ccd where time resolved rss spectra are recorded using a focal plane mask together with the spectrograph slit. spectra of the target and sky are obtained in the rectangular region, which shuffled down across the frame transfer boundary at the end of a short exposure. 2 time-resolved studies of polars the first high-speed observations conducted with salt were mostly of eclipsing magnetic cataclysmic variable stars (mcvs), specifically polars. these systems emit the bulk of their luminosity (from x-rays to the optical / ir) from small accretion regions near the magnetic pole(s) of a strongly magnetized (∼101 to 102 mg) white dwarf. the dominant emission components in such systems includes hard (> 10 kev) thermal bremsstrahlung x-rays from an accretion shock just above the white dwarf surface. a fraction of this radiation is reprocessed in the white dwarfs photosphere into softer x-rays, uv and optical radiation. in some systems, ballistic ”blobs” of accreting material bury into the white dwarf’s photosphere, thermalize and contribute to the soft x-ray/extreme uv component. shock cooling occurs through polarized optical/ir cyclotron emission. polars therefore present a wealth of multi-wavelength observational opportunities, which are being exploiting with salt. 2.1 light curves of eclipsing polars in eclipsing polars, the size, structure and location of accretion regions can be determined from high time resolution (sub-second) photometry. since 2005 observations have been carried out with salticam to obtain continuum eclipse curves with unprecedented time resolution and signal to noise. an example is for the eclipsing system fl cet, with a 1.45 h orbital period (o’donoghue et al. 2006). an eclipse egress light curve, taken with salticam in ”slotmode” with 112 ms data sampling, is shown in fig. 4. the intensity steps are due to the progressive re-appearance of two accretion hot-spots near the magnetic poles, which take ∼1.5 s to be uncovered. these data have been fitted with a model based on the likely masses of the component stars, as derived from the eclipse parameters and orbital period, plus the orbital inclination of the system (i) and the co-latitudes of the accretion spots (β1 and β2). figure 4: salticam eclipse egress light curve of fl cet, showing the re-appearance of two accretion spots. other polars are also being systematically observed in order to derive similar quality light curves from which to determine the accretion geometries. an example is shown below for the system hy eri, which has been observed in both high and low accretion states. figure 5: eclipse light curves of hy eri, at high (upper curve) and low (lower curve) accretion states. 2.2 spectra of eclipsing polars since 2011 the eclipse light curve program has been extended to spectroscopic investigations of the emission lines of eclipsing polars, using rss in time-resolved spectroscopy mode. the aim of these observations is to probe the mass transfer flow from the secondary star to the white dwarf, specifically where it changes from a 135 david a. h. buckley ballistic to a magnetically controlled trajectory. the geometry of this so-called ”threading region” varies from system to system and is dependent on the accretion rate and magnetic field topology. previous attempts at probing the nature of the threading region (e.g. hakala et al. 2002, bridge et al. 2004) have often been frustrated by the problem of unambiguously determining the accretion stream structures from the 1-d data derived from broad band eclipse light curves. these new salt observations are aimed to address this issue and resolve details of the mass transfer stream through high time resolution emission line spectroscopy. such data will lead to stream eclipse light curves as a function of wavelength and hence radial velocity, enabling the structure of the mass transfer stream to be determined. doppler tomographic techniques will be applied to these data in order to attempt to map the positions of the line emitting regions. recently a new ”inside-out” doppler tomography technique has been devised (e.g. see kotze & potter; these proceedings). eventually we plan to undertake phase resolved spectropolarimetric observations of these systems, using rss (when the polarimetric modes are re-activated in 2014). these observations will probe the phase dependent cyclotron emission, giving us an independent view of the polarized emission arising from the accretion spots. during low states it may also be possible to study the polarization of the zeeman lines from the white dwarf and the accretion halo. figure 6: an example of time resolved (10s) trailed spectroscopy of the eclipsing polar hu aqr. 2.3 qpos in magnetic cvs in 2012 a program to study quasi-periodic oscillations (qpos) in magnetic cvs was instigated at saao, mostly using the new electron multiplication ccd (emccd) cameras (shoc) on the saao 1.9-m telescope. the aim is to search for and characterise the nature of the optical qpos, which have typical periods of a few seconds. the commonly held understanding is that they are due to plasma oscillations in the magnetically confined accretion columns which constrain the accreted plasma before it cools and settles onto the surface of the magnetic white dwarf. the qpo frequency is thought to be related to the physical parameters inside the column (or sub-columns), particularly the shock height, accretion rate and magnetic field strength. this work will hopefully lead to a better understanding of the physics underlying the qpos, taking the study from being purely phenomenological to one based on quantitative physical parameters. in conjunction with this observational program, we expect that the experimental results from a parallel program of simulating accreting columns in the laboratory will yield further relevant insights (e.g. michaut et al. 2012). the project involves new technology detectors, namely emccds, which have both good qe and fast timing capability (e.g. ∼10 ms). these detectors are superior to those used in the first studies of polar qpos (e.g. larsson 1989, and references therein), which relied on pmt (photomultiplier tube)-based instruments. high time resolution photometry of mcvs have recently been carried out at salt, utilizing the newly installed (in 2012) berkeley visible image tube (bvit), a photon counting instrument using a microchannel plate (mcp) detectors (buckley et al., 2010, welsh et al. 2011). while not of particularly high qe, compared to a ccd, bvit is capable of time tagging photons to ∼50 ns, and is ideal for studying extremely fast variability (e.g. pulsars). short (typically ∼1 h) bvit observations have been made of several mcvs, but the only positive detection to date has been for the intermediate polar, ex hya, where a ∼100 s qpo (and possibly its harmonic), were detected (fig. 7). figure 7: a salt observations of ex hya using bvit which revealed a ∼100/50 s qpo. 136 observing cvs and lmxbs with salt:updates and recent results 3 salt spectroscopy of xss j12270-4859 figure 8: salt spectra of xss j12270-4859, averaged over ∼1 h on 4 different nights. the enigmatic fermi gamma ray source, xss j122704859 (e.g. bonnet-bidaud et al. 2012, and references therein), was the subject of a saao-eso observing campaign in march/april 2012, utilizing the ntt, salt and several smaller saao telescopes. spectroscopy was carried out with salt over 6 nights, contemporaneous with ntt spectroscopy. in addition, high speed u-band photometry and jhk imaging was also undertaken. xss j12270-4859 was first suggested to be an intermediate polar (butters et al. 2008), but this was not confirmed from follow-up suzaku and xmm observations. the system exhibits x-ray flares and dips, somewhat reminiscent of exo0748-676, but without any phase dependency and no spectral changes. the x-ray spectrum is also characterized by a weakly absorbed power law, with no cut-off up to 100 kev and no fe lines. the system has been suggested to be similar to a microquasar (e.g. cyg x-3, but it is not as variable in x-rays) or possibly a millisecond pulsar (e.g. psr 1023+0038), but no spin period has been detected to date. the salt spectra were obtained during ∼1 h tracks, with 300 s repeat exposures. they show significant changes, often over the timescale of an exposure, sometimes exhibiting weak emission lines of the balmer series, hei and heii, and other time showing balmer and other absorption features, as can be seen in fig. 8. radial velocity variations of the combined spectra indicate a possible ∼7 h orbital period, which is under investigation. the high speed photometry also shows variations on a similar timescale, but with periods of enhanced flickering possibly occurring at the same orbital phase. 4 future plans in 2014 it is planned to extend the salt observational work on cvs and related objects to include time resolved polarimetric imaging and spectropolarimetry. polarized continuum flux variations will be used to probe the cyclotron emission, for example using the stokes imaging technique (see potter et al. 2004). this will be combined with doppler and roche tomography and eclipse mapping to assist in building detailed models of mcvs, and may include using relevant particle or mhd codes. following the launch of the indian x-ray satellite, astrosat (in 2014), it is planned to begin coordinated observations using salt to undertake contemporaneous optical observations of mcvs, lmxbs and related objects. astrosat will have both, hard and soft x-ray instruments, the former with excellent sensitivity and energy range, plus a uv imaging telescope. adding such data will allow for a comprehensive multiwavelength analysis of these systems. 137 david a. h. buckley acknowledgement the work summarized in this paper represents a number of collaborative salt projects with various colleagues, including jean marc bonnet-bidaud, hannes breytenbach, domitilla de martino, marissa kotze, darragh o’donoghue and stephen potter, to name a few. the research of the author is supported by nrf rated researcher and incentive funding awards. the observations presented were carried out using the southern african large telescope. references [1] bonnet-bidaud, j.m., et al.: 2012, mem. s.a. it., 83, 742. [2] bridge, c.m., et al., 2004, mnras, 351, 1423. doi:10.1111/j.1365-2966.2004.07884.x [3] buckley, d.a.h., swart, g.p & meiring, j.g., 2006, proc. spie, 6267, 62670z. doi:10.1117/12.673750 [4] buckley, d.a.h., et al., 2010, poprc. spie, 7735, 77359-1. [5] butters, o.w., et al., 2008, a&a, 487, 271. [6] kotze, e. & potter, s.b., 2013, these proceedings. [7] larsson, s., 1989, a&a, 217, 146. [8] michaud, c., et al., 2012, mem. s.a. it., vol. 83, 665. [9] o’donoghue, d., et al., 2006, mnras, 372, 151. doi:10.1111/j.1365-2966.2006.10834.x [10] potter, s.b., et al., 2004, mnras, 348, 316. doi:10.1111/j.1365-2966.2004.07379.x [11] welsh, b., et al., 2011, iau symp. 285, 99. discussion paula szkody: can you give more information on the qpo in ex hya? is it seen in the lines or continuum? david buckley: the qpo was detected using salt and bvit with a broad-band r-filter, so includes both continuum and hα. the characteristic period was ∼100 s, or more likely, its harmonic. 138 http://dx.doi.org/10.1111/j.1365-2966.2004.07884.x http://dx.doi.org/10.1117/12.673750 http://dx.doi.org/10.1111/j.1365-2966.2006.10834.x http://dx.doi.org/10.1111/j.1365-2966.2004.07379.x introduction high speed photometry with salticam time resolved spectroscopy with rss time-resolved studies of polars light curves of eclipsing polars spectra of eclipsing polars qpos in magnetic cvs salt spectroscopy of xss j12270-4859 future plans 49 acta polytechnica ctu proceedings 1(1): 49–55, 2014 49 doi: 10.14311/app.2014.01.0049 planck 2013 cosmology results: a review josé alberto rubiño-mart́ın1,2 on behalf of the planck collaboration 1instituto de astrofisica de canarias, e-38200 la laguna, tenerife, spain 2departamento de astrofisica, universidad de la laguna, e-38206 la laguna, tenerife, spain corresponding author: jalberto@iac.es abstract this talk presents an overview of the cosmological results derived from the first 15.5 months of observations of the esa’s planck mission. these cosmological results are mainly based on the planck measurements of the cosmic microwave background (cmb) temperature and lensing-potential power spectra, although we also briefly discuss other aspects of the planck data, as the statistical characterization of the reconstructed cmb maps, or the constraints on cosmological parameters using the number counts of galaxy clusters detected by means of the sunyaev-zeldovich effect in the planck maps. all these results are described in detail in a series of papers released by esa and the planck collaboration in march 2013. keywords: cosmology: observations cosmic microwave background cosmological parameters galaxy clusters. 1 introduction planck1 (tauber et al., 2010; planck collaboration i, 2011) is the third generation space mission to measure the anisotropy of the cosmic microwave background (cmb). it observes the sky in nine frequency bands covering 30–857 ghz with high sensitivity and angular resolution from 31’ to 5’. the low frequency instrument (lfi; mandolesi et al., 2010; bersanelli et al., 2010; mennella et al., 2011) covers the 30, 44, and 70 ghz bands with amplifiers cooled to 20 k. the high frequency instrument (hfi; lamarre et al., 2010; planck hfi core team, 2011a) covers the 100, 143, 217, 353, 545, and 857 ghz bands with bolometers cooled to 0.1 k. polarisation is measured in all but the highest two bands (leahy et al., 2010; rosset et al., 2010). a combination of radiative cooling and three mechanical coolers produces the temperatures needed for the detectors and optics (planck collaboration ii, 2011). two data processing centres (dpcs) check and calibrate the data and make maps of the sky (planck hfi core team 2011b; zacchei et al., 2011). planck’s sensitivity, angular resolution, and frequency coverage make it a powerful instrument not only for cosmology, but also for galactic and extra-galactic astrophysics. this talk focuses on the main cosmology results based on planck’s second data release, which covers data acquired in the period 12 august 2009 to 27 november 2010, and will be quoted here as nominal mission maps. all these results and the associated data products are described in a series of papers released in march 2013 (planck collaboration i–xxix, 2013). because the analysis of the polarization data is not yet as mature as the analysis of the temperature data, the planck polarization results are not included in this release. 2 overview of planck 2013 cosmology results the main goal of the planck mission is to determine with great precision the key cosmological parameters describing our universe, using the information encoded in the cmb anisotropies on intermediate and small angular scales over the whole sky (tauber et al., 2010; planck collaboration i, 2013). the data processing pipelines and beam calibrations employed by the lfi and hfi dpcs to create and characterize the nine full-sky maps based on the first 15.5 months of operations are described in planck collaboration ii–ix (2013). these nine maps (see table 1) allow robust reconstruction of the primordial cmb temperature fluctuations over nearly the full sky, as well as new constraints on galactic foregrounds, including thermal dust and line emission from molecular carbon monoxide (planck collaboration xii 2013). four different component separation methods were optimized to produce 1 planck (http://www.esa.int/planck) is a project of the european space agency (esa) with instruments provided by two scientific consortia funded by esa member states (in particular the lead countries france and italy), with contributions from nasa (usa) and telescope reflectors provided by a collaboration between esa and a scientific consortium led and funded by denmark. 49 http://dx.doi.org/10.14311/app.2014.01.0049 josé alberto rubiño-mart́ın a cmb map based on the planck maps alone, i.e., without the addition of any other external data. figure 1 shows the map produced by the smica method. 2.1 the planck cmb likelihood and the angular power spectrum a complete statistical description of the two-point correlation function of the cmb temperature fluctuations measured by planck which accounts for all known relevant uncertainties, is encoded in the planck likelihood (planck collaboration xv 2013). we follow a hybrid approach to construct this likelihood, using a exact approach at large scales (` < 50), and a pseudo-c` approach at small scales (` > 50). this function can be used to derive our estimate of the cmb angular power spectrum (see figure 2), which constitutes an extremely precise measurement over three decades in multipole range, allowing to see seven acoustic peaks. in this plot, our main source of error at ` ≤ 1500 is due to cosmic variance. the measurements are in excellent agreement with previous experiments (wmap, act, spt). we have validated our likelihood through an extensive list of consistency tests, assessing the impact of residual foreground and instrumental uncertainties on the final results. although the polarization data are not released, we have checked that the best-fit λcdm cosmology is in excellent agreement with preliminary planck ee and te polarisation spectra. table 1: planck performance parameters determined from flight data (planck collaboration i 2013). channel νcenter fwhm noise sensitivity [ghz] [arcmin] [µkrj s 1/2] 30 ghz 28.4 33.16 145.4 40 ghz 44.1 28.09 164.8 70 ghz 70.4 13.08 133.9 100 ghz 100 9.59 31.52 143 ghz 143 7.18 10.38 217 ghz 217 4.87 7.45 353 ghz 353 4.70 5.52 545 ghz 545 4.73 2.66 857 ghz 857 4.51 1.33 2.2 lensing by large scale structure on small (arcminute) angular scales, the cmb anisotropies are perturbed by gravitational lensing, primarily sourced by the large scale structure of the universe at relatively high redshifts (peaking at z ∼ 2). planck provides a 25σ detection of this effect (planck collaboration xvii, 2013), allowing also a lensing potential reconstruction at the map level on almost the full sky. the power spectrum of this lensing potential map is used to construct a “lensing” likelihood, which is included in our cosmological analyses. the inclusion of the planck lensing reconstruction can break degeneracies inherent in the temperature data alone, especially the geometric degeneracy in nonflat models, thus providing a strong constraint on spatial curvature without the need of including polarization data. 2.3 cosmological parameters: base model the planck measurements of the cmb temperature and lensing-potential power spectra are well described by the standard spatially-flat six parameter λcdm model with adiabatic scalar perturbations (planck collaboration xvi, 2013). the computed values of the six key parameters within this model are summarized in table 2, together with the values obtained for some derived parameters, as ωλ, ωm, σ8 or the hubble constant h0. the most remarkable result is that, for this cosmology, we find a low value of the hubble constant, which is in tension with recent direct measurements of h0, but is in excellent agreement with constraints from baryon acoustic oscillation surveys (see planck collaboration xvi (2013) for a detailed discussion). we also find a higher value of the matter density, as compared to previous measurements. 2.4 cosmological parameters: beyond the base model multiple one-parameter extensions to the base (sixparameter) model have been discussed in planck collaboration xvi (2013). however, the main result is that there is no compelling evidence for any of these extensions. for example, including curvature, we find that the universe is consistent with spatial flatness to subpercent level precision (ωtot = 1.0010 ± 0.0065, at 95%, using planck +lensing+wp+highl+bao). there is no evidence for additional neutrino-like relativistic particles beyond the three families of neutrinos in the standard model. using bao and cmb data, we find neff = 3.30 ± 0.27 for the effective number of relativistic degrees of freedom, in excellent agreement with the standard value of 3.046, and an upper limit of 0.23 ev for the sum of neutrino masses. in addition, we find no evidence for dynamical dark energy; using bao and cmb data, the dark energy equation of state parameter 50 planck 2013 cosmology results: a review is constrained to be w = −1.13+0.13−0.10. finally, deviations from the simplest inflationary scenario are discussed in planck collaboration xvi and xxii (2013), including tests of specific inflationary models, isocurvature modes, and broken scaleinvariance. as a main result, planck rules out exact scale invariance (ns = 1) at high significance. the deviation of the scalar spectral index from unity is insensitive to the addition of tensor modes and to changes in the matter content of the universe. we find a 95% upper limit of r < 0.11 on the tensor-to-scalar ratio (see figure 3), which allows us to constrain a number of inflationary models. figure 1: the smica method estimates the cmb over about 97% of the sky, with the remaining area replaced with a constrained gaussian realization. figure taken from planck collaboration i (2013). figure 2: cmb temperature angular power spectrum from planck. the shaded area represents cosmic variance. the solid line corresponds to the best-fit six-parameter λcdm model quoted in the text. figure taken from planck collaboration i (2013). 51 josé alberto rubiño-mart́ın table 2: cosmological parameter values for the six-parameter base λcdm model. column 2 give the results for the planck temperature and lensing potential power spectra, while column 3 adds the information from wmap polarization (wp) at low multipoles (` ≤ 23), and other small scale cmb experiments (highl). see planck collaboration xvi 2013 for details. the last four rows are derived parameters. planck + lensing + wp + highl ωbh 2 0.02217 ± 0.00033 0.02218 ± 0.00026 ωch 2 0.1186 ± 0.0031 0.1186 ± 0.0022 100θmc 1.04141 ± 0.00067 1.04144 ± 0.00061 τ 0.089 ± 0.032 0.090+0.013−0.014 ns 0.9635 ± 0.0094 0.9614 ± 0.0063 ln(1010as) 3.085 ± 0.057 3.087 ± 0.024 ωλ 0.693 ± 0.019 0.693 ± 0.013 ωm 0.307 ± 0.019 0.307 ± 0.013 h0 67.9 ± 1.5 67.9 ± 1.0 σ8 0.823 ± 0.018 0.8233 ± 0.0097 figure 3: marginalized joint 68% and 95% confidence level regions for ns and r from planck in combination with other data sets, compared to the theoretical predictions of some inflationary models. taken from planck collaboration xxii (2013). 2.5 gaussianity and statistical isotropy the statistical isotropy and the gaussian nature of the planck cmb maps has been rigorously examined in several papers (planck collaboration xxiii–xxvi, 2013). as a main result, there is no detection of primordial non gaussianity, obtaining the following values for the primordial local, equilateral and orthogonal bispectrum amplitudes: flocalnl = 2.7 ± 5.8, f equil nl = −42 ± 75, and forthonl = −25 ± 39 (all at 68% confidence level). the integrated sachs-wolfe-lensing bispectrum signal is detected at 2–3 sigma level. the planck data also provides stringent new constraints on cosmic strings and other defects (planck collaboration xxv 2013). alternative geometries and non-trivial topologies have also been analyzed (planck collaboration xxvi 2013). deviations from isotropy have been found in planck 52 planck 2013 cosmology results: a review cmb maps, essentially confirming all the low-angle cmb anomalies previously seen in wmap data at similar levels of significance (about 3σ): the low quadrupole amplitude; the quadrupole-octupole alignment; the hemispherical asymmetry; and the “cold spot”. finally, planck notes a further low-angle discrepancy, related to the deficit of power in the angular power spectrum in the multipole range 20 ≤ ` ≤ 40, which could be connected with some of the previous features (e.g. the hemispherical asymmetry or the low observed variance). 2.6 galaxy clusters planck maps can also be used to detect and study galaxy clusters by means of their sunyaev-zeldovich (sz) effect (sunyaev & zeldovich, 1972). early planck results on clusters were published in planck collaboration viii-xii (2011). in march 2013, we have released the planck catalogue of sz sources (psz1) based on the nominal mission maps (planck collaboration xxix 2013). psz1 contains 1227 entries, out of which 861 are confirmed galaxy clusters, and it constitutes the largest all-sky cluster catalogue to date (see figure 4). in planck collaboration xx (2013), we present constraints on cosmological parameters using the number counts (n(z)) for a well-defined and 100% reliable subsample of 189 galaxy clusters from psz1. using a xray calibrated mass-sz flux relation, and assuming a 20% bias between the x-ray determined mass and the true mass of a cluster, we derive constraints on the matter power spectrum amplitude and the matter density parameter, obtaining σ8(ωm/0.27) 0.3 = 0.78±0.01, with one-dimensional ranges σ8 = 0.77 ± 0.02 and ωm = 0.29 ± 0.02. these results are in excellent agreement with the cosmological constraints from the angular power spectrum of the thermal sz emission. in planck collaboration xxi (2013), we construct the first all-sky maps of the thermal sz emission, and compare their power spectrum to theoretical models, finding σ8(ωm/0.27) 0.3 = 0.784 ± 0.016. however, these results favour somewhat low values of σ8 and ωm as compared to the cmb based analysis (see table 2). this tension can be eliminated by relaxing the assumption on the mass bias between x-ray and true mass up to 45%, although such a high value is difficult to reconcile with the known observed properties of galaxy clusters. there is another interesting possibility to reconcile these values, which consists in allowing for a component of massive neutrinos with a total mass (for the sum of the three species) of the order to ∑ mν ≈ 0.22 ev (see marulli et al. 2011 for a review of how cosmological observations can be affected by the inclusion of neutrino masses). these possibilities will be further investigated in future releases. figure 4: distribution of the 1227 planck sz clusters and candidates across the sky, shown in mollweide projection in galactic coordinates. the masked regions in black color (point-sources, magellanic clouds and galactic emission) cover a total area of 16.3%, and are not used by the sz cluster-finder algorithms. taken from planck collaboration xxix (2013). 3 conclusions the main conclusions of the planck 2013 release are: • the planck data are in remarkable agreement with a flat six-parameter λcdm model; • we detect with high significance lensing of the cmb by intervening matter, providing evidence for dark energy from the cmb alone; • the best-fit model contains a weaker cosmological constant (∼ 69%), and more cold dark matter (∼ 26%) than previously estimated; and we firmly establish a deviation from scale invariance for primordial matter perturbations; • we find a low value of the hubble constant, in tension with other astrophysical measurements; • there is no compelling evidence for extensions of the base model; • we find no evidence for significant deviations from gaussianity in the statistics of cmb anisotropies; however, we confirm the anomalies at large angular scales first detected by wmap and find a deficit of power at low-multipoles with respect to our best-fit model. • the planck sz cluster measurements are in tension with the cmb results; although this conclusion is limited by our knowledge of the observablemass relations, it might provide hints for physics in the neutrino sector. 53 josé alberto rubiño-mart́ın the next (2014) release of planck data, which will include the full-mission maps and the polarization information, will provide a significant improvement in data quality and the level of systematic error control, helping in answering these open questions. acknowledgement the development of planck has been supported by: esa; cnes and cnrs/insu-in2p3-inp (france); asi, cnr, and inaf (italy); nasa and doe (usa); stfc and uksa (uk); csic, micinn, ja and res (spain); tekes, aof and csc (finland); dlr and mpg (germany); csa (canada); dtu space (denmark); ser/sso (switzerland); rcn (norway); sfi (ireland); fct/mctes (portugal); and prace (eu). a description of the planck collaboration and a list of its members, including the technical or scientific activities in which they have been involved, can be found at http://www.sciops.esa.int/index.php?project=planck. jar-m is grateful 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[42] rosset, c. et al. 2010, a&a, 520, a13 [43] sunyaev, r. a. & zeldovich, y. b. 1972, comments on astrophysics and space physics, 4, 173 [44] tauber, j. a. et al. 2010, a&a, 520, a1 [45] zacchei et al. 2011, a&a, 536, a5 54 planck 2013 cosmology results: a review discussion marco regis: can the large scale anomalies be due to galactic foregrounds ? jose alberto rubiño-martin: we have performed a large number of tests using the four cmb maps produced by the different component separation algorithms, and the consistency of the results strongly favors a cosmological origin for the anomalies. in addition, the agreement between planck and wmap rules out a possible explanation based on systematic effects. 55 introduction overview of planck 2013 cosmology results the planck cmb likelihood and the angular power spectrum lensing by large scale structure cosmological parameters: base model cosmological parameters: beyond the base model gaussianity and statistical isotropy galaxy clusters conclusions 55 acta polytechnica ctu proceedings 2(1): 55–59, 2015 55 doi: 10.14311/app.2015.02.0055 cataclysmic variables from sdss: a review and a look forward to lsst p. szkody1 1department of astronomy, university of washington, usa corresponding author: szkody@astro.washington.edu abstract the past and current projects of the sloan digital sky survey are reviewed in the context of applicability and results for cataclysmic variables. ongoing and future time domain surveys that will have impact on the field are also briefly discussed. keywords: cataclysmic variables dwarf novae intermediate polars polars optical photometry spectroscopy. 1 introduction cataclysmic variables (cvs) have been discovered in a variety of ways, with the largest numbers of new systems coming from survey work. the mode of operation and the limiting magnitude of each survey determines the different types of cvs that are discovered. the palomar-green survey (green et al. 1986) used blue color to find objects that were brighter than 16th magnitude, hence it found many bright novalike systems. x-ray surveys such as rosat (voges et al. 1999) identified those systems with high x-ray flux, hence many polars and intermediate polars were found. the hamburg survey (hagen et al. 1995) searched for emission line objects down to about 18th magntitude, finding sw sex stars, intermediate polars and long period dwarf novae. the sloan digital sky survey (sdss) used both photometry (to 22nd magnitude) and spectroscopy (to 19th magnitude) to identify all types of cvs, especially those with low accretion rates and orbital periods below the gap (szkody et al. 2011). current and future surveys are concentrating on variability as a means of identification and as a result are turning up many dwarf novae. the combination of all of these results will ultimately end in a better understanding of the total space density of all types of cvs. 2 sdss the sdss project (york et al. 2000) began taking data in 2000 and has continued to the present time, undergoing changing modes of operation and target selection. the past results on cvs from this survey, as well as current and future plans are summarized below. 2.1 sdss i,ii legacy survey sdss i started with a goal to photometrically survey the entire north galactic cap, as well as obtain spectra of a selected subset of objects within what is termed the legacy survey. in addition to the north cap, a 2.5 degree wide stripe centered on the celestial equator (stripe 82) was also included (see figure 1). the initial plan for completion in 5 years was extended (sdss ii) in order to finish the original footprint so the legacy encompassed the years 2000-2008 and the final entire database was released as data release 7 (dr7; www.sdss.org/dr7/). this database includes imaging data of 230 million objects, using 54 second integrations in 5 filters (ugriz). based on the colors obtained, targets (primarily quasars and galaxies) were selected by a variety of groups, within restrictions of fiber spacing, brightness limits, etc. for 1 hour integration spectra with wavelength coverage from 3800-9200å at a resolution of about 2000. plug plates were than drilled and threaded to accomodate 640 fibers resulting in 1.37 million spectra in dr7, including 225 thousand stars. since the sources receiving spectra were determined by color, and cvs have a broad range of colors (szkody et al. 2003), the main source of cv spectra turned out to be objects taken from quasar loci, which span a broad range of colors outside the main sequence footprint. the resulting computer and eye searches of all the legacy spectra for balmer lines turned up 285 cvs which included 30 polars, 6 ips and 9 systems containing pulsating white dwars. a list of these sources can be found in szkody et al. (2011) as well as on the web with links to the spectra (http://www.astro.washington.edu/users/szkody/cvs/). from 2000-2011 extensive followup on more than 300 55 http://dx.doi.org/10.14311/app.2015.02.0055 p. szkody nights was conducted on this list by many observers using apo, la palma, usno, steward, mdm, mmt observatories. as a result of this work, 151 orbital periods were determined that allowed specific classification of these objects and how they fit into population models. gänsicke et al. (2009) used 126 of the periods known at that time to reach major conclusions that enforced the magnetic braking model and confirmed the population synthesis predictions of howell et al. (2001): i.e. the majority of the disk accreting systems exist below the period gap and a period spike appears at the minimum period. the sdss results showed that the discrepancies in the past were primarily due to selection effects that favored the discovery of bright, long period systems, while sdss was able to uncover the larger population of faint sources. however, the percentages under the gap are slightly less and the period spike occurs at a slightly longer period than predicted so adjustments to the angular momentum losses are needed (knigge 2011). the fainter magntitude limit reached by sdss also revealed large numbers of accreting pulsators and low accretion rate polars (larps). of the 16 known accreting pulsators, 9 were found in the legacy survey. all of these showed broad balmer absorption lines surrounding the emission, providing a clear optical signature of the white dwarf. followup optical and uv results on the set of accreting pulsators has led to three major results. the first is that the instability strip is much wider than h atmosphere pulsating white dwarfs (zz ceti) (szkody et al. 2010). arras et al. (2006) attribute this to the existence of a he instability strip as well as hydrogen. because many of the accreting white dwarfs in this wide instability strip are observed not to pulsate, followup long term observations were conducted. this led to the second result that objects can stop pulsating, usually after an outburst, but sometimes when no outburst has occurred! e.g. sdss0745+45 (eq lyn, mukadam et al. 2013). third, in followup observations with hst, it was discovered that the pulsation that was present after outburst in v455 and appeared in the emission lines, not the continuum, presenting problems for a physical mechanism (szkody et al. 2013). of the 9 known larps, 7 were found in sdss. these objects have prominent humps due to cyclotron harmonics at high fields and low optical depth (wickramasinghe & ferrario 2000). finding larps in sdss data is not easy as the dependence of the harmonics on field strength moves them into different regions of color space. schmidt et al. (2005) calculated the sdss color ranges for various fields but there has as yet been no systematic search to find all the candidates. they estimate these could be a major contribution to the magnetic white dwarf population. several groups are using the sdss database to further understand the properties of the known cvs and to find further ones. a large hst program led by boris gänsicke is obtaining uv spectra of 40 cvs (including 12 discovered by sdss) to characterize the temperature of the white dwarf and the mass accretion rates over a variety of orbital periods and compositions. when combined with gaia distances, the masses of these white dwarfs will also be determined, leading to improved understanding of the evolution of close binaries. carter et al. (2013) discovered 29 new cvs in doing spectroscopic followup of am cvn candidates selected by color in the photometric database. breedt and gänsicke (2011) are obtaining spectra of the faintest cvs in sdss that are being discovered by crts (see section below). 2.2 sdss ii: segue and sn besides the legacy extension, sdss ii (2005-2008) contained the sloan extension for galactic understanding and exploration (segue) and a supernova survey. these results are contained in dr8. the main goal of segue was to explore the milky way, including its structure, history, kinematics, evolution and dark matter by mapping the positions, velocities, composition and temperatures of 240,000 stars (see figure 1 for the coverage compared to legacy). while this survey mainly targetted white dwarfs and giants, several cvs were included as their colors matched those of white dwarfs. the sn survey involved repeated imaging of the same region of the sky (stripe 82) during 3 months of each year. the repeat imaging over 275 deg2 with about 20 measurements on each object identified hundreds of new transient sources (sako et al. 2008). while the main goal was identifying sn, other types of variable stars were also found. the analysis of all the light curves of variable objects is ongoing and may include some cvs. a catalog of 13,051 variable sources brighter than g=20.5 from 1998-2007 stripe 82 data is in sesar et al. (2007). bhatti et al. (2010) provide a catalog of light curves for 221,842 point sources for half of the entire stripe 82 data. 2.3 sdss iii: segue-2, boss, apogee and marvels sdss iii (2008-2014) continues the segue project and adds 3 new enterprises. segue-2 obtained spectra of 119,000 stars with a concentration on the stellar halo with distances of 10-60 kpc. these data appeared in dr8 while dr9 updated the stellar parameters and added catalogs. the baryon oscillation spectroscopic survey (boss) updated the sdss fibers to 1000 and is producing spectra of many galaxies and quasars. while the target selection is not as optimal as it was in the 56 cataclysmic variables from sdss: a review and a look forward to lsst legacy survey, there are a few cvs which emerge in the spectra. dr9 was the first public release of boss spectra, while dr10 contains the latest data. the apo galactic evlution experiment (apogee) uses ir spectra to observe red giants throughout the galaxy. the first data appear in dr10. finally, the multiobject apo radial velocity exoplanet large-area survey (marvels) is monitoring the radial velocities of 11,000 bright stars to look for planets. 2.4 sdss iv: the future plans are underway to extend sdss from 2014-2020 with a continuation that involves apogee-2, eboss and manga. apogee-2 will continue the milky way exploration using apo and extend to the south with a 2.5m telescope at las campanas. eboss will continue boss but add 2 segments of interest to cvs: a time-domain spectroscopic survey (tdss ) that will obtain spectra of 100,000 variable sources and the spectroscopic identification of erosita sources (spiders). mapping nearby galaxies at apo (manga) will obtain spatially resolved spectra of 10,000 nearby galaxies. 3 current surveys several surveys are now ongoing and searching for objects that vary. the catalina real-time transient survey (crts; drake et al. 2009) consists of 3 telescopes: a 1.5m on mt. lemmon, a 0.7m on catalina and a 0.5m at siding springs. at the time of this meeting, 1022 potential cvs with outburst magnitudes brighter than 17 were posted on the web page nesssi.cacr.caltech.edu/catalina/brightcv.html. this page has a column that denotes whether the object is in the sdss photometric database and provides a direct link to sdss. since most of the objects were found at outburst, the quiescent magnitudes tend to be very faint (20-22nd mag) and therefore difficult to followup spectroscopically. several groups are now following up on the crts sources (woudt et al. 2012; thorstensen & skinner 2012). as noted above, elme breedt is also leading a project using gemini and other large telescopes to categorize the faintest crts sources that have 5 color photometry in sdss. the panoramic survey telescope & rapid response system (panstarrs1) is using an 1.8m telescope on haleakala, hawaii to complete a 2010-2013 northern sky transient survey which observes the available sky several times a month (tonry et al. 2012). one of the 12 key projects involves variables and explosive transients. a data release is planned for 2014 and a second telescope is under development. the palomar transient factory (ptf) involves the palomar 48 in schmidt telescope to image the sky from 2009-2014 to a magnitude of 21 on timescales from minutes to years to find new transients and variables, including cvs (rau et al. 2009). recent improvements (called iptf) have led to pipeline products that provides candidates within 30 minutes for spectroscopic followup. the followup spectroscopy is accomplished with the palomar 1.5m and other telescopes and some of that is available in wiserep (yaron & gal-yam 2012). skymapper is a robotic 1.35m survey telescope in siding springs, australia that is imaging the entire southern sky 36 times over 5 years in a series of 6 filters to 22nd mag that will provide spectral types of stars as well as variability information (keller et al. 2007). the data will be made public through the virtual observatory. in addition to these single, wide-field telescope surveys, there are 2 all-sky surveys which are using 2-4 cameras to image the sky. one is the all sky automated survey (asas) which images the entire sky to 14th mag in v and i bands from las campanas, chile and haleakala, maui (pojmanski 1997). the other is the mobile astronomical system of telescope-robots (master) which images to 19th mag at sites from russia and argentina (lipunov et al. 2010). 4 the future: lsst the future for investigation of variability lies with the large synoptic survey telescope (lsst). this 8.4m telescope situated in chile will image 18,000 deg2 of sky about 1000 times over 10 years (2020-2030). it uses 6 filters and will reach r = 24.5 mag on single nights and 27.5 on co-added images. the survey will produce alerts within minutes of observation as well as long term catalogs that will be made public. details may be found in the online science book (http://www.lsst.org/lsst/scibook). this survey will go several magnitudes fainter than sdss and should be able to find the population of period bounce systems that are predicted by models, as well as find unusual long term variability such as found by honeycutt et al. (2003) during their long term monitoring with roboscope. however, the difficulty lies in planning spectroscopic followup for 24-25 mag objects, which will require a lot of observing time on the largest telescopes available. time series photometry for short period, low amplitude variables will still be possible, but the problems of smart classification to pick out interesting variables from the multitude each night and the ensuing spectral confirmation remain to be solved. 5 conclusions the sdss has provided a significant database of cvs including a consistent set of medium resolution spectra for 285 systems and a photometric database that 57 p. szkody likely contains many more at faint magnitudes down to 22. due to the fainter magnitude limit compared to previous surveys, followup observations resulted in a large change in the observed orbital period distribution of cvs that has resolved some discrepancies in close binary evolution. current and future surveys rely on discoveries based primarily on variability, and so uncover large numbers of dwarf novae. as these surveys push further into the galactic plane and to fainter magnitudes with larger telescopes, the true space density of cvs and the distribution among types will become better known. however, the detailed information that comes from spectroscopy will be difficult to obtain for the faintest systems. acknowledgement the work with sdss data took place by many individuals over many years, starting with the sdss collaboration and extending to the public. support for the cv part of the research was provided by nsf grants ast 97-30792, ast 0607840 and ast 1008734 and nasa grant hst-go-12870.07a. funding for sdssiii has been provided by the alfred p. sloan foundation, the participating institutions, the national science foundation, and the u.s. department of energy office of science. the sdss-iii web site is http://www.sdss3.org/. sdss-iii is managed by the astrophysical research consortium for the participating institutions of the sdss-iii collaboration including the university of arizona, the brazilian participation group, brookhaven national laboratory, carnegie mellon university, university of florida, the french participation group, the german participation group, harvard university, the instituto de astrofisica de canarias, the michigan state/notre dame/jina participation group, johns hopkins university, lawrence berkeley national laboratory, max planck institute for astrophysics, max planck institute for extraterrestrial physics, new mexico state university, new york university, ohio state university, pennsylvania state university, university of portsmouth, princeton university, the spanish participation group, university of tokyo, university of utah, vanderbilt university, university of virginia, university of washington, and yale university. references [1] arras, p., townsley, d. m. & bildsten, l.: 2006, apj 643, l119. doi:10.1086/505178 [2] bhatti, w. a. et al. : 2010, apjs 186, 233. doi:10.1088/0067-0049/186/2/233 [3] breedt, e. & gänsicke, b. t.: 2011, aspcs 447, 203. [4] carter, p. j. et al.: 2013, mnras 429, 2143. doi:10.1093/mnras/sts485 [5] drake, a. j. et al.: 2009, apj 696, 870. doi:10.1088/0004-637x/696/1/870 [6] gänsicke, b. t. et al.: 2009, mnras 397, 2170. doi:10.1111/j.1365-2966.2009.15126.x [7] green, r. f., schmidt, m. & liebert, j.: 1986, apjs 61, 305. [8] hagen, h. j. et al.: 1995, aaps 111, 195. [9] honeycutt, r. k. et al.: 2003, aspcs 292, 329. [10] howell, s. b., nelson, l. a. & rappaport, s.: 2001, apj 550, 897. doi:10.1086/319776 [11] keller, s., bessell, m., schmidt, b. & francis, p.: 2007, aspcs 364, 177. [12] knigge, c.: 2011, aspcs 447, 3. [13] lipunov, v. et al.: 2010, adast 30l. [14] mukadam, a. s. et al.: 2013, aj 146, 54. [15] pojmanski, g.: 1997, aa 47, 467. [16] rau, a. et al.: 2009, pasp 121, 1334. [17] sako, m. et al.: 2008, aj 135, 348. [18] schmidt, g. d. et al.: 2005, apj 630, 1037. [19] sesar, b. et al.: 2007, apj 134, 2236. [20] szkody, p. et al.: 2003, aj 126, 1499. [21] szkody, p. et al.: 2010, apj 710, 64. doi:10.1088/0004-637x/710/1/64 [22] szkody, p. et al.: 2011, aj 142, 181. [23] szkody, p. et al.: 2013, apj 775, 66. doi:10.1088/0004-637x/775/1/66 [24] thorstensen, j. r. & skinner, j. n.: 2012, aj 144, 81. [25] tonry, j. l. et al.: 2012, apj 750, 99. doi:10.1088/0004-637x/750/2/99 [26] voges, w. et al.: 1999, aap 349, 389. [27] wickramasingeh, d. t. & ferrario, l.: 2000, pasp 112, 873. [28] woudt, p. et al.: 2012, mnras 421, 2414. [29] yaron, o. & gal-yam, a.: 2012, pasp 124, 668. doi:10.1086/666656 [30] york, d. g. et al.: 2000, aj 120, 1579. 58 http://dx.doi.org/10.1086/505178 http://dx.doi.org/10.1088/0067-0049/186/2/233 http://dx.doi.org/10.1093/mnras/sts485 http://dx.doi.org/10.1088/0004-637x/696/1/870 http://dx.doi.org/10.1111/j.1365-2966.2009.15126.x http://dx.doi.org/10.1086/319776 http://dx.doi.org/10.1088/0004-637x/710/1/64 http://dx.doi.org/10.1088/0004-637x/775/1/66 http://dx.doi.org/10.1088/0004-637x/750/2/99 http://dx.doi.org/10.1086/666656 cataclysmic variables from sdss: a review and a look forward to lsst figure 1: the sky coverage of the legacy and segue surveys in galactic (left) and celestial (right) coordinates, from www.sdss.org/dr7. discussion linda schmidtobreick: wouldn’t you expect the pulsating white dwarfs in accreting systems to be hotter than single white dwarfs? due to accretion, you would get a thin hot layer that influences the measured temperature but not necessarily the pulsation. paula szkody: yes, most accreting white dwarfs are indeed observed to be hotter than single white dwarfs. however, there are several parameters that could affect the pulsations (instability strip) in these accreting white dwarfs besides the temperature. the accreting ones are spun up by the accretion and the composition of the atmosphere is different due to the mass transfer from the secondary. right now, we don’t have enough data to distinguish which of these parameters are determining whether an accreting white dwarf will pulsate or not. 59 introduction sdss sdss i,ii legacy survey sdss ii: segue and sn sdss iii: segue-2, boss, apogee and marvels sdss iv: the future current surveys the future: lsst conclusions acta polytechnica ctu proceedings doi:10.14311/app.2016.3.0019 acta polytechnica ctu proceedings 3:19–24, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app the effect of support plate on drilling-induced delamination navid zarif karimia, ∗, hossein heidaryb, parnian kianfarc, mahmud hasanic, giangiacomo minaka a university of bologna, via fontanelle 40, forli, italy b university of tafresh, tehran road 1, tafresh, iran c amirkabir university of technology, hafez 424, tehran, iran ∗ corresponding author: navid.zarif@unibo.it abstract. delamination is considered as a major problem in drilling of composite materials, which degrades the mechanical properties of these materials. the thrust force exerted by the drill is considered as the major cause of delamination; and one practical approach to reduce delamination is to use a back-up plate under the specimen. in this paper, the effect of exit support plate on delamination in twist drilling of glass fiber reinforced composites is studied. firstly, two analytical models based on linear fracture mechanics and elastic bending theory of plates are described to find critical thrust forces at the beginning of crack growth for drilling with and without back-up plate. secondly, two series of experiments are carried out on glass fiber reinforced composites to determine quantitatively the effect of drilling parameters on the amount of delamination. experimental findings verify a large reduction in the amount of delaminated area when a back-up plate is placed under the specimen. keywords: drilling, composite materials, delamination, support plate. 1. introduction glass fiber reinforced plastics (gfrps) have been extensively employed in many engineering applications because of their outstanding advantages over other materials. the machining and specially drilling of gprps have become very important due to the need for assembling of subcomponents made from these materials. but, composite materials are difficult to machine due to some particular characteristics of them like non-homogeneous, anisotropic and abrasive fibers. this causes significant damages in drilling process such as matrix cracking, fiber breakage, fuzzing and thermal degradation. among the damages caused by drilling, delamination is considered as one the most crucial, which results in lowering of strength against fatigue and subsequently reducing the longterm performance of composite structures [1, 2]. several techniques have been suggested to reduce delamination in drilling process that each has its own advantages and limitations. one of the frequently used methods is to place a support plate under the specimen. during supported drilling, in contrast to unsupported drilling, the specimen is not free to bend due to constraints imposed by back-up plate. in drilling with back-up plate, the thrust force increases suddenly and severely as the chisel edge begins to penetrate into the specimen and it continues until the full penetration of cutting lips. at this point actual material removal takes place and the force reaches its steady state value, until delamination occurs. in this stage, delamination causes the force to decrease stepwise, each step associates to a crack opening. unlike supported drilling, during drilling without back-up plate, the force increases gradually at the beginning of process. this is mainly because the real feed rate is much lower than the adjusted feed due to relative movement of drill bit and workpiece. as the drill bit approaches the last uncut plies, the stress due to the thrust force exceeds its critical value and all the uncut material is burst open [3]. until now, many researchers have attempted to reduce drilling induced delamination. hocheng and dharan used linear elastic fracture mechanics to determine the critical thrust force for twist drilling as push-out delamination begins to propagate [4]. hocheng and tsao proposed some analytical models for different drill bits including candle stick drill, saw drill, core drill and step drill to determine the thrust force at the beginning of delamination growth [5, 6]. based on these models, they realized the delamination can be reduced by lowering the thrust force or distributing the force outward from the center [7]. in a similar work, they studied the effect of exit back-up plate on drilling-induced delamination when using a slot drill bit and a core drill bit [8]. they showed that the critical thrust force is increased when using a back-up plate, hence delamination is less likely to happen. many efforts have been made to reduce the thrust force exerted by drill bit on the specimen, while few researchers studied the effect of support plate on thrust force and associated delamination. this paper investigates the effect of back-up plate on drilling-induced delamination both analytically and experimentally. 19 http://dx.doi.org/10.14311/app.2016.3.0019 http://ojs.cvut.cz/ojs/index.php/app n. zarif karimi, h. heidary, p. kianfar et al. acta polytechnica ctu proceedings 2. experimental procedure 2.1. delamination measurement chen presented an index to evaluate the amount of delamination called conventional delamination factor [9]. however, this factor is not proper because the crack size does not represent the damage magnitude appropriately and also this procedure does not indicate the damage area. davim et al. suggested a superior approach to determine the delamination factor called the adjusted delamination factor fda which is expressed as eq. (1) [10]. the first part of eq. (1) shows the size of the crack contribution and the second part shows the damage area contribution. fda = α dmax d0 + β amax a0 (1) where the coefficients α and β can be calculated as below: β = ad a0 − amax α = 1 − β (2) where d0 is the nominal diameter of the hole, dmax is the maximum diameter of the damage hole, a0 is the area related to the nominal hole, amax is the area related to the maximum diameter of the delamination zone and ad is the delaminated area. 2.2. materials and tools the composite plates were produced by hand lay-up method with araldite ly556 epoxy resin reinforced with 60 % e-glass unidirectional fiber. the density of epoxy resin and glass fibers were 1.12 g/cm3 and 292 g/m2, respectively. tensile strength and tensile modules of resin were 80 mpa and 2.7 gpa, respectively, and for glass fibers were 2150 mpa and 74 gpa, respectively. the composite laminates had 16 plies and a thickness of 5 mm. the holes were drilled at the center of plates by standard hss twist drills 10 mm. during unsupported drilling, a back-up plate of 5 mm thick was positioned under the specimens. an appropriate clamping system was used to fix the specimens and back-up plate in the drilling machine, shown in figure 1. the specimens and back-up plate were fixed in position by tightening four screws at the corners. 2.3. plan of experiments in the present study, taguchi method was used to design drilling experiments. three parameters, feed rate, cutting speed and drill point angle were selected and three levels for each parameter were suggested based on our preliminary researches [11–14]. after determination of parameters and correspond levels, a proper orthogonal array need to be selected. the orthogonal array chosen was the l9. after conducting experiments, the signal to noise ratio (sn) needs to be calculated for each experiment to determine the effect of each parameter. according figure 1. experimental setup. to this method, the optimization is done by using three signal to noise ratios; smaller is better, larger is better and nominal is better. notice that, regardless of what approach is taken, always higher value of the signal to noise ratio is better. in this investigation, the delamination factor needs to be minimized; hence, smaller is better definition of the signal to noise ratio was used. sni = −10 log( ni∑ i=1 y2i ni ) (3) where i is the experiment number and ni is the number of trials for experiment i [15]. 3. results and discussion 3.1. theoretical analysis of critical thrust force the energy balance equation at the onset of delamination propagation is: dud = dw − du (4) in which du is the infinitesimal strain energy, dw is the infinitesimal work done by the thrust force fth and drill displacement of dx and dud is the infinitesimal strain energy absorbed by crack growth which are as following: dw = fth · dx dud = gic · da (5) where da is the change in the delamination area and gic is the critical strain energy release rate in mode i, which is assumed to be constant according to saghizadeh and dharan [16]. 3.1.1. critical thrust force in drilling without support plate figure 2 depicts the schematic of delamination in unsupported drilling. in figure 2, fth is the thrust force exerted by a twist drill at the center of plate, x is the displacement of drill, h is the thickness of 20 vol. 3/2016 the effect of support plate on drilling-induced delamination figure 2. the schematic of delamination for unsupported drilling. specimen, h is the uncut depth under tool, and a is the radius of crack (delamination). in this model, two assumption are considered; the isotropic behavior and pure bending of the laminate. according to the classical plate theory, for a circular plate with clamped edges and concentrated force the amount of deflection can be expressed as [17]: w(r) = fth 16πm [2r2 ln r a + (a2 − r2)] (6) which is written in polar coordinate system (r is radius). m = eh3/(12(1 − ϑ2)) is flexural rigidity of the plate, e is modulus of young, and ϑ is ratio of poisson. the stored strain energy, the work done and the strain energy absorbed by crack growth are expressed as below equations, respectively. u = f 2tha 2 32πm du = f 2tha 16πm da (7) w = fthw(0) = f 2tha 2 16πm dw = f 2tha 8πm da (8) ud = gic · a = gicπa2 dud = gic 2πada (9) now it is possible to calculate the critical thrust force at the onset of crack propagation by replacing eqs. (7), (8) and (9) in the energy balance eq. (4) as shown below: fth = π √ 32gicm (10) in order to prevent delamination, the thrust force exerted by the drill bit on the specimen which is related to the material properties and the uncut thickness should not go over this value. 3.1.2. critical thrust force in drilling with support plate figure 3 depicts the schematic of delamination in drilling with back-up plate. in figure 3, fb is the thrust force with back-up, r is the upward reaction force from the back-up plate, b is the radius of the applied ring force of back-up and c is the drill radius, x is the displacement. for an edge clamped circular plate under the action of a concentrated force in center and upward circular force, the amount of deflection can be expressed as: w1(r) = fb 16πm [2r2 ln r a + (a2 − r2)]− − r 8πm [(r2 + b2) ln b a + 1 2 (1 − b2 a2 )(a2 + r2)] 0 6 r 6 b (11) w2(r) = fb 16πm [2r2 ln r a + (a2 − r2)]− − r 8πm [(r2 + b2) ln r a + 1 2 (1 + b2 a2 )(a2 − r2)] b 6 r 6 a (12) using boundary conditions, w(b) = 0, we get: r = fb [b2 ln b a + ( b 2−a2 2 )] [2b2 ln b a + ( a4−b42a2 )] (13) by replacing r into the plate deflection equations, eq. (11) and eq. (12), the total stored strain energy and work done by external forces can be derived as follow: u = π[ ∫ b 0 [m( ∂2w1 ∂r2 + 1 r ∂w1 ∂r )2]rdr+ + ∫ a b [m( ∂2w2 ∂r2 + 1 r ∂w2 ∂r )2]rdr] du = ∂u ∂a da (14) w = fbw(0) = fb [ fba 2 16πm − r 8πm (b2 ln b a + + ( a2 − b2 2 ))] dw = ( f 2ba 8πm − rfb 8πm a2 − b2 a )da (15) finally, by calculating dud from eq. (14) and replacing dud and dw in the energy balance equation, eq. (4), we get: fb = π √ 32gicm b a (16) where a and b are defined as follow: 21 n. zarif karimi, h. heidary, p. kianfar et al. acta polytechnica ctu proceedings figure 3. the schematic of delamination for supported drilling. b = [a4 − b4 − 8a2b2 ln b a ]× × [64b4a4 ln(a2b2) − 128a4b4 ln b ln a + 16a6b2 ln b a + + 16a2b6 ln a b + a8 + b8 − 2a4b4] (17) a = 3(a12 − b12) + 72a10b2 ln b a + 192a4b8 ln a ln b+ + 3072a6b6 ln a2 + b b2 + a + 768a4b8 ln a2 + b b2 + a − − 288a8b4 ln a ln b − 704a6b6 ln b + 192a2b10 ln a ln b+ + 73a4b4(a4 − b4) + 32a2b2(a8 − b8) + 64b4a8 ln b a + + 144a6b6 ln b a + 56a2b10 ln b a + 352a6b6 ln a2b2− − 96a2b10 ln a2b2 + 144a8b4 ln a2b2 + 1024a6b6 ln b3 a3 + + 256a4b8 ln b3 a3 + 768a4b8 ln a ln b − 16b12 ln a2b2+ + 32b12 ln a ln b − 384a4b8 ln a2b2 (18) 3.2. experimental results the graphs of thrust force and the scanned images of delamination in drilling with and without support plate at a feed rate of 0.025 mm/rev, spindle speed of 1600 rpm and drill point angle of 130 ◦ are shown in figure 4. it is believed that the thrust force applied by the drill on the specimen causes delamination and it occurs when the thrust force goes beyond its critical value. however, the graphs of the force in figure 4 clearly indicate this is only true for drilling with a support plate. during unsupported drilling the force is lower than in supported drilling, whereas the delamination is more extensive, which suggests that in unsupported drilling a different mechanism is in play, and other factors must be considered. these factors can be the dynamics of the workpiece, and the way in which the force is applied. the measured adjusted delamination factor (fda) and the corresponding signal to noise ratio (sn) determined using taguchi method smaller is better for figure 4. the graphs of thrust force and the scanned images of delamination in drilling with and without support plate. figure 5. comparison of fda in drilling with and without back-up. the two cases of drilling, supported and unsupported, are reported in table 1.symbols i and ii in table 1 represent two measured values for fda. the t-test on the mean values of adjusted delamination factor demonstrates significant differences between the means in these two cases (t0 = 8.00 > t0.005,34 = 2.73). to assist in the practical interpretation of this experiment, it is helpful to construct a graph of average responses at each experiment condition described in table 1. this graph is shown in figure 5. from figure 5, a significant difference in adjusted delamination factor for supported drilling and unsupported drilling can be observed so that in all drilling tests, the average value of delamination is reduced when applying a back-up plate. this reduction is in the range of 8-27 % for different drilling conditions. this reduction is mainly due to the reduction in matrix crack growth as a result of upward reaction of back-up plate. the values of sn data for the adjusted delamination factor for unsupported and supported drilling are given in table 1 and shown in figure 6. for unsupported drilling, the most important parame22 vol. 3/2016 the effect of support plate on drilling-induced delamination test no. parameters unsupported supported feed rate(mm/rev) speed (rpm) angle (◦) fda(1) fda(2) s/n fda(1) fda(2) s/n 1 0.025 800 90 1.41 1.37 -2.86 1.04 1.05 -0.34 2 0.025 1250 118 1.25 1.28 -2.04 1.07 1.06 -0.55 3 0.025 1600 130 1.30 1.26 -2.14 1.11 1.07 -0.75 4 0.05 800 118 1.49 1.52 -3.55 1.09 1.09 -0.87 5 0.05 1250 130 1.20 1.21 -1.62 1.10 1.11 -1.25 6 0.05 1600 90 1.33 1.28 -2.31 1.16 1.15 -0.75 7 0.1 800 130 1.49 1.58 -3.73 1.15 1.14 -1.40 8 0.1 1250 90 1.32 1.30 -2.35 1.18 1.17 -1.17 9 0.1 1600 118 1.36 1.33 -2.57 1.22 1.25 -1.83 table 1. l9 orthogonal array of taguchi and experimental results. figure 6. effect of process parameters for (a) unsupported drilling, (b) supported drilling. ters affecting the delamination factor are the spindle speed followed by feed rate. the optimum process parameters on the delamination are obtained as feed rate at level 1 (0.025 mm/rev), spindle speed at level 2 (1250 rpm) and drill point angle at level 3 (130 ◦) when drilling without back-up plate. however, when the back-up plate is applied, the effect of feed rate on delamination increases so that it becomes the most important parameter followed by spindle speed, figure 6 (b). the cause of this phenomenon can be attributed to the cutting mechanism and the relative movement of the tool and specimen. it is observed that the optimal value of the feed rate should be kept at low level in order to minimize the delamination factor. the optimum process parameters on the delamination are obtained as feed rate at level 1 (0.025 mm/rev), spindle speed at level 1 (800 rpm) and drill point angle at level 3 (130 ◦) for supported drilling. 4. conclusions in this paper, the effect of exit support plate on delamination in drilling of glass fiber reinforced composites is studied. a comprehensive analysis of the critical thrust force in drilling with and without back-up plate is presented based on three theories i.e. energy conservation theory, elastic bending theory and linear elastic fracture mechanics. the experimental results confirm the effects of delamination reduction when using a back-up plate under the specimen. this reduction is in the range of 8-27 % for different drilling conditions. this reduction is attributed to the crack growth suppression by upward reaction of back-up plate. references [1] r. m. jones. mechanics of composite materials. crc press, 1998. [2] d. liu, y. tang, w. l. cong. a review of mechanical drilling for composite laminatesc. composite structures 94(4):1265–1279, 2012. [3] e. capello. workpiece damping and its effect on delamination damage in drilling thin composite laminates. journal of materials processing technology 148(1):186–95, 2004. [4] h. ho-cheng, c. k. h. dharan. delamination during drilling in composite laminates. journal of manufacturing science and engineering 112(1):236–9, 1990. [5] h. hocheng, c. c. tsao. comprehensive analysis of delamination in drilling of composite materials with various drill bits. journal of materials processing technology 140(1):335–9, 2003. [6] h. hocheng, c. c. tsao. effects of special drill bits on drilling-induced delamination of composite materials. international journal of machine tools and manufacture 46(1):1403–16, 2006. [7] h. hocheng, c. c. tsao. the path towards delamination-free drilling of composite materials. journal of materials processing technology 167(1):1251– 64, 2005. doi:10.1016/j.ijmachtools.2005.10.004. [8] h. hocheng, c. c. tsao. effects of exit back-up on delamination in drilling composite materials using a saw drill and a core drill. international journal of machine tools and manufacture 45(1):1261–70, 2005. 23 http://dx.doi.org/10.1016/j.ijmachtools.2005.10.004 n. zarif karimi, h. heidary, p. kianfar et al. acta polytechnica ctu proceedings [9] w.-c. chen. some experimental investigations in the drilling of carbon fiber-reinforced plastic (cfrp) composite laminates. international journal of machine tools and manufacture 37(1):1097–108, 1997. [10] j. p. davim, j. c. rubio, a. m. abrao. a novel approach based on digital image analysis to evaluate the delamination factor after drilling composite laminates. composites science and technology 67(1):1939–45, 2007. [11] n. zarif karimi, h. heidary, m. ahmadi. residual tensile strength monitoring of drilled composite materials by acoustic emission. materials and design 40(1):229–36, 2012. [12] n. zarif karimi, h. heidary, g. minak, m. ahmadi. effect of the drilling process on the compression behavior of glass/epoxy laminates. composite structures 98(1):59– 68, 2013. doi:10.1016/j.compstruct.2012.10.044. [13] h. heidary, n. z. karimi, m. ahmadi, et al. clustering of acoustic emission signals collected during drilling process of composite materials using unsupervised classifiers. journal of composite materials 49(5):559–71, 2015. doi:10.1177/0021998314521258. [14] n. zarif karimi, g. minak, p. kianfar. analysis of damage mechanisms in drilling of composite materials by acoustic emission. composite structures 131(1):107– 14, 2015. doi:10.1016/j.compstruct.2015.04.025. [15] r. k. roy. a primer on the taguchi method, second edition. society of manufacturing engineers, 2010. [16] h. saghizadeh, c. k. h. dharan. delamination fracture toughness of graphite and aramid epoxy composites. journal of engineering materials and technology 108(1):290–5, 1986. [17] j. n. reddy. mechanics of laminated composite plates and shells: theory and analysis. crc press, 2004. 24 http://dx.doi.org/10.1016/j.compstruct.2012.10.044 http://dx.doi.org/10.1177/0021998314521258 http://dx.doi.org/10.1016/j.compstruct.2015.04.025 acta polytechnica ctu proceedings 3:19–24, 2016 1 introduction 2 experimental procedure 2.1 delamination measurement 2.2 materials and tools 2.3 plan of experiments 3 results and discussion 3.1 theoretical analysis of critical thrust force 3.1.1 critical thrust force in drilling without support plate 3.1.2 critical thrust force in drilling with support plate 3.2 experimental results 4 conclusions references 26 acta polytechnica ctu proceedings 2(1): 26–32, 2015 26 doi: 10.14311/app.2015.02.0026 the space density of magnetic and non-magnetic cataclysmic variables, and implications for cv evolution m. l. pretorius 1department of physics, university of oxford, denys wilkinson building, keble road, oxford ox1 3rh, united kingdom 2previous address: school of physics and astronomy, university of southampton, highfield, southampton so17 1bj corresponding author: retha.pretorius@astro.ox.ac.uk abstract we present constraints on the space densities of non-magnetic and magnetic cataclysmic variables, and discuss some implications for models of the evolution of cvs. the high predicted non-magnetic cv space density is only consistent with observations if the majority of these systems are extremely faint in x-rays. the data cannot rule out the very simple model where long-period ips evolve into polars and account for the entire short-period polar population. the fraction of wds that are strongly magnetic is not significantly higher for cv primaries than for isolated wds. finally, the space density of ips is high enough to explain the bright, hard x-ray source population seen in the galactic centre. keywords: cataclysmic variables dwarf novae nova-likes intermediate polars polars x-rays. 1 introduction there are many remaining uncertainties in the theory of cataclysmic variable (cv) formation and evolution, as well as several serious discrepancies between predictions and the properties of the observed cv population (e.g. patterson 1998; pretorius, knigge & kolb 2007a; pretorius & knigge 2008a,b; knigge, baraffe & patterson 2011). in order to constrain evolution models, we require more precise observational constraints on the properties of the galactic cv populations. a fundamental parameter predicted by evolution theory, that should be more easily measured than most properties of the intrinsic cv population, is the space density, ρ. a few specific, important open questions concerning the formation and evolution of cvs are: (i) is the large predicted population of non-magnetic cvs at short orbital period consistent with the current observed cv sample? (ii) is there an evolutionary relationship between ips and polars? (iii) can the intrinsic fraction of mcvs be reconciled with the incidence of magnetic wds in the isolated wd population? (iv) do mcvs dominate the total galactic x-ray source populations above lx ∼ 1031 ergs−1? these questions can be addressed empirically, with reliable measurements of the space densities of the different populations of cvs. uncertainty in ρ measurements is in part caused by statistical errors, resulting from uncertain distances and small number statistics. however, the dominant source of uncertainty is most likely systematic errors caused by selection effects. fig. 1 shows some reported measurements (differing by several orders of magnitude for non-magnetic cvs). selection effects are most easily accounted for in samples with simple, well-defined selection criteria. in the absence of a useful volume-limited cv sample, a purely flux-limited sample is the most suitable for measuring ρ. whereas optical cv samples always include selection criteria based on, e.g., colour or variability, there are x-ray selected cv samples that are purely flux-limited. all active cvs show x-ray emission generated in the accretion flow. furthermore, mcvs are luminous x-ray sources, while the correlation between the ratio of optical to x-ray flux and the optical luminosity of non-magnetic cvs, implies that an x-ray flux limit does not introduce as strong a bias against shortperiod cvs as an optical flux limit (e.g. van teeseling et al. 1996). here we use 2 x-ray surveys, the rosat bright survey (rbs; e.g. schwope et al. 2002), and the rosat north ecliptic pole (nep) survey (e.g. henry et al. 2006) to construct x-ray flux-limited cv samples. we then provide robust observational constraints on the space densities of both magnetic and non-magnetic cvs, by carefully considering the uncertainties involved. 26 http://dx.doi.org/10.14311/app.2015.02.0026 the space density of magnetic and non-magnetic cataclysmic variables, and implications for cv evolution figure 1: some previously reported measurements of the space densities of different cv populations. we provide additional background on the questions listed above in section 2, present the measurements in section 3, discuss the implications of the results in section 4, and finally list the conclusions in section 5. 2 context 2.1 missing non-magnetic cvs it is not clear whether the present-day observed nonmagnetic cv population is inconsistent with theoretical expectations. population synthesis models predict that only ' 1 percent of all cvs are long-period systems (see e.g. kolb 1993). the the vast majority of cvs should therefore be intrinsically faint. pretorius et al. (2007a) and pretorius & knigge (2008b) used a specific model of kolb (1993), together with models of the outburst properties and seds of cvs, to show that, although observed cv samples are strongly biased against shortperiod systems, a currently undetected faint cv population cannot dominate the overall population to the extent predicted by this particular population synthesis model. knigge et al. (2011) used the properties of cv donor stars to conclude that the aml rate is lower above the gap and higher below the gap than predicted by the standard model. this leads to larger predicted factions of both period bouncers and long-period systems. whether this is consistent with observed cv samples is not yet known. that a large faint population of cvs exists is now clear from observations (gansicke et al. 2009; patterson 2011). however, whether observations have truly revealed a population as large as predicted remains to be seen. some predicted values of the non-magnetic cv space density are as high as 2 × 10−4 pc−3 (de kool 1992; kolb 1993); most observational estimates are much lower, but values ranging from ≤ 5×10−7pc−3 to ρ ∼ 10−4pc−3 have been reported. perhaps the most straight forward test of these predictions is to compare them to the measured space density of the galactic short-period, non-magnetic cv population. 2.2 evolutionary relationship between ips and polars in most regards, the formation and evolution of magnetic and non-magnetic cvs is believed to be similar. both types of cvs form via common envelope (ce) evolution, evolve initially from long to short porb as a result of angular momentum loss (aml), and eventually experience period bounce, when the thermal timescale of the donor becomes longer than the mass transfer time-scale. the main proposed difference between the evolution of mcvs and non-magnetic cvs actually affects only the polars, where magnetic braking (mb) is thought to be suppressed because of the very strong wd magnetic field (e.g. li & wickramasinghe 1998; townsley & gansicke 2009). the porb distributions of magnetic and non-magnetic cvs are broadly consistent with these ideas. considering polars and ips together, their porb distribution is very similar to that of non-magnetic cvs, showing a period gap in the range 2 hr <∼porb <∼3 hr, as well as a period minimum at around porb ' 80 min (see fig. 2). it has long been recognized that most ips are found above the period gap and most polars below (fig. 2)., which immediately suggests that ips may evolve into polars (e.g. chanmugam & ray 1984). physically this makes sense, since smaller orbital separation and lower ṁ (besides large magnetic field strength) favour synchronization. 27 m. l. pretorius figure 2: the orbital period distribution of all cvs (black), and mcvs (green). cumulative distributions are also shown for polars (red) and ips (blue). almost all ips are found at long porb, while most polars are short-porb systems. periods from ritter & kolb (2003). mb drives much higher mass-transfer rates above the period gap than gravitational radiation (gr) does below the gap; therefore, it is plausible that many accreting magnetic wds may only become synchronized after crossing the period gap. the main problem with this scenario is that the magnetic fields of the wds in ips (bip <∼10mg) are systematically weaker than those of the wds in polars (bpolar ∼ 10−100mg). there are several possible solutions to this problem. perhaps the most simple (in terms of binary star evolution) is that the high accretion rates in ips partially “bury” the wd magnetic fields, so that the observationally inferred field strengths for these systems are systematically low (cumming 2002). it is also possible that the short-period polar population is dominated by systems born below the period gap. of course this does not explain why we observe so few short-period ips, although it may be that longperiod ips become unobservable once they reach short periods (see patterson 1994; wickramasinghe, wu & ferrario 1991). a way to investigate the relationship between ips and polars is to compare their respective space densities. for example, if all long-period ips evolve into short-period polars, and all short-period polars form out of of long-period ips, then their space densities should be proportional to the evolutionary time-scale associated with these two evolutionary phases. in this particular example, we would predict that ρpolar/ρip ' τgr/τmb >> 1. 2.3 intrinsic fraction of mcvs magnetic systems make up ' 20% of the known cv population (ritter & kolb 2003). taken at face value, this is a surprisingly high fraction, considering that the strong magnetic fields characteristic of ips and polars (b >∼10 6 g) are found in only ' 10% of isolated wds (e.g. kulebi et al. 2009). if these numbers really represent the intrinsic incidence of magnetism amongst cvs and single wds, the difference between them would have important implications, namely that either strong magnetic fields favour the production of cvs, or that some aspect of pre-cv evolution favours the production of strong magnetic fields (see e.g. tout et al. 2008). however, it is in reality not yet clear that magnetism is more common in cv primaries than in isolated wds. the main problem is that the observed fraction of magnetic systems amongst known cvs is almost certainly affected by serious selection biases. for example, since mcvs are known to be more luminous in x-rays than non-magnetic cvs, they are likely to be over-represented in x-ray-selected samples. conversely, polars, in particular, are relatively faint in the optical band (since they do not contain optically bright accretion disks), so they are likely to be under-represented in optically-selected samples. given that the overall cv sample is a highly heterogeneous mixture of x-ray, opticaland variability-selected sub-samples (which also typically lack clear flux limits), it is very difficult to know how the observed fraction of mcvs relates to the intrinsic fraction of magnetic wds in cvs. 28 the space density of magnetic and non-magnetic cataclysmic variables, and implications for cv evolution 2.4 galactic x-ray source populations several studies have attempted to determine the makeup and luminosity function of galactic x-ray source populations in different environments, including the milky way as a whole, the galactic centre, the galactic ridge, and globular clusters. remarkably, in all of these environments, mcvs have been proposed as the dominant population of x-ray sources above lx >∼10 31ergs−1. in most of these studies, identifying the observed xray sources with distinct populations is very uncertain, since few of the sources have either optical counterparts or properties that allow for clear classification. the classification of observed sources rely mainly on x-ray luminosities and hardness, and statistical comparisons of observed and expected number counts. the local space densities of the relevant physical populations are likely the most important aspect of these comparisons. in effect, the relevant question is whether the extrapolation of the local space density to the environment being investigated can account for the number of sources observed there. in the case of mcvs, such extrapolations are difficult, mainly because the local space densities are quite poorly known. 3 calculating space densities 3.1 the flux-limited samples the rbs covers |b| > 30◦ to fx >∼10 −12erg cm−2s−1, and includes 16 non-magnetic cvs, and 30 mcvs (6 ips and 24 polars). the nep covers 81 sq.deg. to fx >∼10 −14. only 4 cvs where detected in the nep, all of them non-magnetic. the samples are presented in pretorius et al. (2007b), pretorius & knigge (2012), and pretorius et al. (2013). 3.2 the method we use the 1/vmax method (e.g. stobie et al. 1989) together with a monte carlo simulation designed to sample the full parameter space allowed by the data, as described in pretorius et al. (2007b) and pretorius & knigge (2012). we tested the method to verify that it gives reliable error estimates, and also considered various possible systematic biases (pretorius & knigge 2012; pretorius et al. 2013). 3.3 results 3.3.1 probability distribution functions the distributions of mid-plane ρ values, normalized to give probability distribution functions, from the simulations are shown in fig. 3. the best-estimate midplane space densities are 4+6−2 × 10 −6 pc−3 for nonmagnetic cvs, and 8+4−2 × 10 −7 pc−3 for mcvs. for the 2 classes of mcvs, we find 3+2−1 × 10 −7 pc−3 for ips and 5+3−2 × 10 −7 pc−3 for polars. 3.3.2 upper limits on ρ of undetected populations the ρ estimates assume that the detected populations are representative of the intrinsic populations, in that they contain at least 1 of the faintest systems present in the intrinsic populations. it is possible that even large populations of sources at the faint ends of the luminosity functions can go completely undetected in flux-limited surveys. we performed additional monte carlo simulations to place limits on the sizes of faint populations of cvs that could escape detection in the surveys we have used (see pretorius & knigge 2012; pretorius et al. 2013). fig. 4 shows the maximum allowed ρ as a function of lx , separately for possible undetected nonmagnetic, polar and ip populations. specifically, if ρn−m = 2×10−4 pc−3 (at the high end of the predicted range), we require that the majority of non-magnetic cvs have lx <∼4 × 10 28 erg s−1. a population of undetected polars with a space density as high as 5× the measured ρpolar must have lx <∼10 30 ergs−1. a hidden population of ips can only have ρ = 5 × ρip if it consists of systems with x-ray luminosities fainter than 5 × 1030 ergs−1. 4 discussion 4.1 missing non-magnetic cvs we discuss our measured ρn−m, as well as the upper limit, in the context of the predicted (i) large total space density of non-magnetic cvs, and (ii) large predicted fraction of normal short-period cvs and period bouncers. population synthesis models predict that only a few percent of all cvs are above the period gap (kolb 1993 finds less than 1%, while knigge et al. 2011 predict 3%). although we find that long-period systems account for just over 50% of our total ρn−m, the data cannot rule out these theoretical predictions. for example, using the knigge et al. (2011) fraction of longperiod systems, and assuming that we have not significantly under-estimated the space density of longperiod cvs, the space density of short-period cvs is ' 2 × 10−6 pc−3(97/3) ' 6 × 10−5 pc−3. the upper limit on ρn−m from section 3.3.2 then implies that a short-period cv population of this size could escape detection in the two surveys, as long as these systems have lx <∼8 × 10 28 erg s−1 (for the simple case of a hypothetical single-lx population of faint, undetected cvs). the predicted ρ = 2 × 10−4 pc−3 would require that most cvs have x-ray luminosities below lx = 4 × 1028 erg s−1. 29 m. l. pretorius figure 3: the ρ distributions for non-magnetic cvs (left-hand panel) and mcvs (right-hand panel) resulting from our simulations. solid lines in both panels mark the modes, medians, and means of the distributions; dashed lines show 1-σ intervals. the probability distribution functions shown in the inset are for the whole mcv sample, polars alone (red), and ips alone (blue). reproduced from pretorius & knigge (2012) and pretorius, knigge & schwope (2013). figure 4: the upper limit on ρ as a function of x-ray luminosity for undetected populations of cvs. the left-hand panel is for non-magnetic cvs, and the right-hand panel for ips (blue) and polars (red). in the lefthand panel, the 2 upper, fine histograms show the corresponding results for the rbs (middle) and nep (top) surveys alone. note that the assumed x-ray spectra of non-magnetic cvs, polars, and ips are different, hence the different slopes. reproduced from pretorius & knigge (2012) and pretorius et al. (2013). 4.2 evolutionary relationship between ips and polars if one assumes that long-period ips are the only progenitors of short-period polars, and that all ips become polars after crossing the period gap, then the ratio of the space densities of long-porb ips and short-porb polars (ρip,lp and ρpolar,sp) will simply reflect the ratio of evolutionary time-scales. the observed logarithm of this ratio is log (ρpolar,sp/ρip,lp) = 0.32 ± 0.36. if the evolution of long-period ips is driven by mb, while that of short-period polars is driven by gr alone, the evolutionary time-scale of short-period polars should be >∼5× that of long-period ips (e.g. knigge et al. 2011). this is completely consistent with the ratio of the inferred space densities. in fact, at 2-σ, the uncertainties are large enough to allow both ratios exceeding 10 and ratios below one. this means that, with the best existing space density estimates for polars and ips, we cannot strongly constrain the evolutionary relationship between these two classes. nevertheless, it is interesting to note that the simplest possible model, in which short-period polars derive from long-period ips and all ips become polars, is not ruled out by their observed space densities. 30 the space density of magnetic and non-magnetic cataclysmic variables, and implications for cv evolution 4.3 intrinsic fraction of mcvs combining the space density estimates of magnetic and non-magnetic cvs to estimate the intrinsic fraction of mcvs, we find log(fmcv ) = −0.80+0.27−0.36. this is consistent, within the considerable uncertainties, with the fraction of isolated wds that are strongly magnetic . furthermore, it is possible that the x-ray-selected cv sample is less complete for non-magnetic cvs than for mcvs. therefore, the incidence of magnetism is not obviously higher amongst cv primaries compared to isolated wds. 4.4 galactic x-ray source populations we consider if it is plausible that ips dominate x-ray source populations above lx ' 1031ergs−1, taking the galactic centre as an example. the deep chandra survey of muno et al. (2009) includes ' 9000 sources down to lx ' 1031 ergs−1, in an area of ' 10−3deg2. approximating the volume covered by the survey as a sphere of radius 150 pc, the space density of x-ray sources in the galactic centre is ρx,gc ∼ 6×10−4 pc−3, compared to the local ip space density of ρip ∼ 3 × 10−7 pc−3. the stellar space density in the galactic centre is ' 1600× higher than in the solar neighborhood. thus these densities are consistent, and we conclude that most of the x-ray sources seen in the galactic centre can indeed be explained as ips. 5 conclusions assumming that the cv samples from the rbs and nep surveys are representative of the intrinsic populations (in the sense that we detected at least one system at the faintest ends of the luminosity functions of those populations), we find ρn−m = 4 +6 −2 × 10 −6 pc−3 and ρmcv = 8 +4 −2×10 −7 pc−3 (ρpolar = 5 +3 −2×10 −7 pc−3 and ρip = 3 +2 −1 × 10 −7 pc−3). the data allow for more than half of non-magnetic cvs having 28.7 < log(lx/erg s −1) < 29.7, and being undetected. however, to reach ρn−m = 2 × 10−4 pc−3 (at the high end of the predicted range), the data requires that the majority of non-magnetic cvs have lx <∼4 × 10 28 erg s−1. the ratio of the space densities of short-period polars to long-period ips is consistent with the simple hypothesis that long-period ips evolve into short-period polars, giving rise to the entire short-period polar population. existing data cannot rule out that strongly magnetic wds have the same incidence amongst cvs as in the field. taking into account the difference in stellar density, the measured local space density of ips can account for the number of bright x-ray sources detected in the galactic centre. acknowledgement i thank the organizes for a successful meeting, and for inviting me to present this review. references [1] chanmugam g., ray a.: 1984, apj, 285, 252 [2] cumming a.: 2002, mnras, 333, 589 doi:10.1046/j.1365-8711.2002.05434.x [3] de kool m.: 1992, a&a, 261, 188 [4] gänsicke b. t., et al.: 2009, mnras, 397, 2170 doi:10.1111/j.1365-2966.2009.15126.x [5] henry j.p., et al.: 2006, apjs, 162, 304 doi:10.1086/498749 [6] knigge c.: 2006, mnras, 373, 484 doi:10.1111/j.1365-2966.2006.11096.x [7] knigge c., baraffe i., patterson j.: 2011, apjs, 194, 28 doi:10.1088/0067-0049/194/2/28 [8] kolb u.: 1993, a&a, 271, 149 [9] külebi b., et al.: 2009, a&a, 506, 1341 [10] li j., wickramasinghe d. t.: 1998, mnras, 300, 718 [11] muno m. p., et al.: 2009, apjs, 181, 110 [12] patterson j.: 1984, apjs, 54, 443 [13] patterson j.: 1994, pasp, 106, 209 [14] patterson j.: 1998, pasp, 110, 1132 [15] patterson j.: 2011, mnras, 411,2695 doi:10.1111/j.1365-2966.2010.17881.x [16] pretorius m. l., knigge c.: 2012, mnras, 419, 1442 doi:10.1111/j.1365-2966.2011.19801.x [17] pretorius m. l., knigge c.: 2008a, mnras, 385, 1471 [18] pretorius m. l., knigge c.: 2008b, mnras, 385, 1485 [19] pretorius m. l., knigge c., kolb u.: 2007a, mnras, 374, 1495 [20] pretorius, m. l., knigge, c., schwope, a. d.: 2013, mnras, 432, 570 doi:10.1093/mnras/stt499 [21] pretorius m. l., et al.: 2007b, mnras, 382, 1279 doi:10.1111/j.1365-2966.2007.12461.x 31 http://dx.doi.org/10.1046/j.1365-8711.2002.05434.x http://dx.doi.org/10.1111/j.1365-2966.2009.15126.x http://dx.doi.org/10.1086/498749 http://dx.doi.org/10.1111/j.1365-2966.2006.11096.x http://dx.doi.org/10.1088/0067-0049/194/2/28 http://dx.doi.org/10.1111/j.1365-2966.2010.17881.x http://dx.doi.org/10.1111/j.1365-2966.2011.19801.x http://dx.doi.org/10.1093/mnras/stt499 http://dx.doi.org/10.1111/j.1365-2966.2007.12461.x m. l. pretorius [22] ritter h., kolb u.: 2003, a&a, 404, 301 [23] schwope a.d., et al.: 2002, a&a, 396, 895 [24] tout c. a., et al.: 2008, mnras, 387, 897 doi:10.1111/j.1365-2966.2008.13291.x [25] townsley d. m., gänsicke b. t.: 2009, apj, 693, 1007 doi:10.1088/0004-637x/693/1/1007 [26] wickramasinghe d. t., wu k., ferrario l.: 1991, mnras, 249, 460 doi:10.1093/mnras/249.3.460 [27] warner b.: 1987, mnras, 227, 23 [28] van teeseling a., beuermann k., verbunt f.: 1996, a&a, 315, 467 32 http://dx.doi.org/10.1111/j.1365-2966.2008.13291.x http://dx.doi.org/10.1088/0004-637x/693/1/1007 http://dx.doi.org/10.1093/mnras/249.3.460 introduction context missing non-magnetic cvs evolutionary relationship between ips and polars intrinsic fraction of mcvs galactic x-ray source populations calculating space densities the flux-limited samples the method results probability distribution functions upper limits on of undetected populations discussion missing non-magnetic cvs evolutionary relationship between ips and polars intrinsic fraction of mcvs galactic x-ray source populations conclusions 84 acta polytechnica ctu proceedings 1(1): 84–89, 2014 84 doi: 10.14311/app.2014.01.0084 agn/starburst connection eleonora sani1 1inaf osservatorio astrofisico di arcetri corresponding author: sani@arcetri.astro.it abstract two main physical processes characterize the activity in the nuclear region of active galaxies: an intense star formation (starburst, sb) and an active galactic nucleus (agn). while the existence of a starburst-agn connection is undisputed, still it is not clear which process dominates the energetic output in both local and high redshift universe. moreover there is no consensus on whether agn fueling is synchronous with star formation or follows it during a post-starburst phase. here i first review how to disentangle the relative sb-agn contribution, then i focus on the physical and geometrical properties of the circumnuclear environment. keywords: agn starburst merging feedback imaging spectroscopy mid-ir x-rays. 1 introduction the issue of agn – star formation (agn–sf) connection in local and distant galaxies is possibly relevant for understanding several processes: from galaxy formation and evolution, and the star formation and metal enrichment history of the universe, to the the origin of the extragalactic background at low and high energies and the origin of nuclear activity in galaxies. it is well known that sf traces the growth of a galaxy in terms of stellar mass, and that galaxies assemble their mass through sb episodes. for example, local ulirgs (ultraluminous infrared galaxies) and high-z smgs (submillimeter galaxies) harbor the large molecular-gas reservoirs that are necessary to switch on sb episodes (tacconi et al. 2010, 2013). on the other hand, agns trace the growth of supermassive black holes (smbhs, with masses mbh > 10 6m�) and quiescent smbhs are believed to dwell in almost all galaxy bulges as agn relics, i.e. the result of a past agn activity (soltan 1982, marconi et al. 2004, merloni & heinz 2008). 1.1 evidence of agn–sf connection a direct link between bulge formation and the growth of the central smbhs has been inferred from tight relations between bh mass mbh and the bulge structural parameters, such as velocity dispersion (ferrarese & merrit 2000, gebhardt et al. 2000, gultekin et al. 2009), luminosity (kormendy & richstone 1995, marconi & hunt 2003, sani et al.2011) and mass (magorrian et al. 1998, mh03, häring & rix 2004). since a sb is a natural consequence of dissipative gaseous processes associated with spheroid formation (e.g. barnes & hernquist 1991) and agn–sf connection dating back to the early universe is implied by these results, a mergerdriven scenario is well suited to reproduce such scaling relations. recently, attention has been drawn to the coexistence of pseudobulges and bhs (nowak et al. 2010, sani et al. 2011). indeed pseudobulges seem to follow their own relation with bhs, hosting less massive bhs than classical bulges (greene, ho, & barth 2008, hu 2009), or they are at least displaced from the scaling relations for classical bulges (sani et al. 2011, kormendy et al. 2011). this may indicate a dichotomy in the formation history of galaxies (major merger vs. secular evolution) and/or that activity in galaxy nuclei can be stochastically driven by local processes such as minor mergers, bar/disk instabilities. in the following i concentrate the discussion on the major-merger scenario, its implications and observational evidences. the reader is anyhow advised about the importance of secular evolution processes, especially in low-luminosity, local seyfert galaxies. 1.2 the major-merger scenario (ulirgs-qso path) the merger between galaxies of approximately the same size is relevant for the cosmological evolution of sbs, quasars and early type galaxies (sanders et al. 1988, silk & rees 1998, fabian 1999, granato et al. 2001, di matteo et al. 2005, hopkins et al. 2008, menci et al. 2008). a schematic view of the main steps in the merging process is given in fig. 1. 84 http://dx.doi.org/10.14311/app.2014.01.0084 agn/starburst connection sfr dmbh/dt figure 1: major-merger scenario. starting from the merger of two spiral galaxies (left side), the merger produces a violent sb and the onset of qso activity (middle) and ends with the formation of a dead spheroid (right side). from hopkins et al. 2008. during first encounters, the mergers between gasrich galaxies (i.e. spirals) drag the gas which fuels both sf and agn activity. thus a violent sb occurs (ulirg-phase) as well as heavily embedded smbh growth (observed as obscured agn). then, the peak of accretion happens when the systems coalesce. at this point the quasar is cleaning the circumnuclear environment thanks to strong winds that blow out the gas, the agn becomes x-ray and optically thin. finally, due to feedback processes both bh growth and further sf are quenched leaving a qso relic in a red galaxy (passive evolution). therefore, obscured qso and feedback processes are key ingredients in the bh– host galaxy co–evolution. in the following, after a brief description of the agn and sf emission, i review the agn-sf observations for both single sources (interesting for their physics/structure) and large samples (to infer statistical properties). 2 the agn vs sf emission an issue in agn–sf studies is to correctly identify agns that represent only a small fraction of sources compared to the population detected at optical and ir wavelengths, and their properties can be elusive. obscured agns can be elusive because their optical emission resembles that of normal galaxies and therefore they cannot be identified by the color-selection techniques used for qsos, or because their spectra do not show any agn signature. their ir emission can be dominated by pah features, typical of star-forming galaxies, instead of continuum emission from hot dust. radio observations reveal the radio-loud population of obscured agns (e.g., radio galaxies), but this comprises only a small fraction, 10%, of the total agn population. a hard x-ray selection fails to detect the luminous narrow-line agns in the numbers predicted by some models of the cosmic x-ray background. polletta and collaborators (2007) show the differences in the spectral energy distributions (seds) of agn and star forming galaxies. the average seds of the three classes show some clear differences. they become increasingly blue in the optical–near-ir (λ < 1 µm), warmer in the ir (i.e. red at λ ' 1−10 µm), and brighter in x-rays in the sequence sf→agn2→agn1. 85 eleonora sani more in detail, most of the diagnostics developed till recently fail in identifying an agn when it is intrinsically faint and/or deeply obscured. as an example of a possible case where the agn is hardly detectable, i consider in fig. 2 an intrinsically faint agn, contributing only 20% to the bolometric source luminosity, and moreover its primary emission is absorbed by gas and dust. only two spectral regions are useful for such an agn detection: the hard x-rays (above 2 kev) and the near-mid infrared between 3 and 10 µm. indeed in the near-ir and optical ranges the dust and gas extinction respectively avoid the agn detection, and at radio wavelengths the sb emission peak dilutes agn emission properties. figure 2: top: simulated seds of a sb (blue) and an agn (black) contributing to 80% and 20% respectively of the lbol. bottom: extinction of the agn emission by gas (green) and dust (red), the horizontal black line indicates an absorption factor of 5. the agn is undetectable in the range 3 µm-2 å due to dust and gas extinction, and at λ ≥ 10 µm for the sb dilution. only the hardest x-rays and 3-10 µm observations reveal the agn. it is thus clear that the diagnostic performed in the x-ray and ir bands can highlight the main agn-sf differences and represent the wavelength rages where we do have the most detailed information respectively. 3 x-ray and ir screenings: imaging and spectroscopy here i briefly describe recent results obtained thanks to x-ray – ir imaging and spectroscopy in two puzzling sources. the final goal is to constrain the physics of ongoing processes and the structure of the nuclear region. 3.1 the case of ngc 6240 thanks to imaging techniques we can simultaneously observe the agn and sf activity in the nuclei of local galaxies (e.g. ngc 1068, ngc 6240, mrk 231, arp 299, circinus...). the angular resolution represents the only limit for this analysis. ngc 6240 is a well known double-merging system 107 mpc far away. here the xray imaging and spectroscopy performed with chandra (komossa et al. 2003) revealed a compton thick double agn, that was confirmed at all wavelengths, plus a sb covering a projected ∼ 2 kpc scale (risaliti et al. 2006, medling et al. 2011, feruglio et al. 2013). hst and spitzer images (bush et al. 2008) show how the thermal emission follows the optical dust obscuration very closely. ngc 6240 is thus an active interacting remnant viewed at the point of nuclei merging where two agns are visible. the results are well consistent with a major merger scenario with the transition from disk galaxies to a spheroid. 3.2 iras 12071-0444: a ct type 2 qso mid-ir (data from irs/spitzer) and x-ray spectroscopy (data from bepposax, chandra, suzaku) allow to investigate the nature and peculiar structure of the nearby (z= 0.128) ulirg (lir = 9 × 1045erg/s) iras 12071-0444. it is one of the few ulirgs optically classified as a type 2 agn and its mid-ir spectrum (in fig. 5–right, nardini et al. 2009) shows the typical features of ongoing sb (traced by pah emission) plus a buried agn (traced by a reddened power-law). however, the marginal emission below 10 kev (fig. 5– left) is consistent with the sole sb component, the typical fe kα line is undetected, and the agn is only seen above 20kev where the much narrower suzaku fov eventually confirms an early tentative detection by bepposax. the x-ray data are well fitted by an absorbed power-law with nh = 3.5 × 1024cm−2 and l2−10 ∼ 7×1044erg/s, i.e. about 10% of the bolometric luminosity is given by the agn intrinsic emission. the result proves iras 12071-0444 as the closest comptonthick quasar 2 yet known. finally, a simple visual comparison with the prototypical ct seyfert ngc 4945 reveals that the reflection efficiency is about a factor of ten lower than for a standard geometry. a peculiar geometry of the absorber is thus required to explain the x-ray/ir data. there are basically two possibilities: (i) a geometrically thick torus, with a small opening angle (i.e. large covering factor). in this case both the direct and reflected components are absorbed. 86 agn/starburst connection agn template sb template figure 3: iras12071-0444 data. left: the spitzer/irs ∼ 5 − 8µm spectrum (green dots) is reproduced (black line) by means of an agn (blue dashed line) and sb (red dot-dashed line) templates. right: suzaku data (blue) well agrees with the chandra/acis spectrum (red) in the 0.5-5 kev range, and beyond 15 kev with the bepposax/pds points (orange). the shaded blue and orange areas in the upper left corner allow to compare suzaku and bepposax/pds fields of view respectively. i also compare iras 12071-0444 with the ngc 4945 x-ray spectrum observed by xmm-newton (green). the ngc 4945 data are scaled to match iras 12071-0444 at 0.5–10 kev. alternatively, (ii) the reflection component is absent due to a small compact cloud along the line of sight (i.e. with a small covering factor). the last picture implies both photoelectric compton absorption without scattering. the resulting intrinsic luminosity exceedes the bolometric one a factor of 3 thus the compact cloud scheme is unphysical. as a probe of the iras 120710444 represents a key evolutionary step of the galaxybh growth. 4 statistical analysis of composite sources by selecting large samples of composite sb+agn sources, it is possible to link the agn–sf connection processes with the system properties. in the following i show recent results concerning ir and x-ray statistical analysis. i ) the availability of a reliable measure of the relative agn/sb contribution to ulirgs (risaliti et al. 2006, sani et al. 2008, nardini et al. 2008) allows a quantitative investigation of the relation between total luminosity and agn contribution. it is known from optical spectroscopy that the fraction of seyfert-like systems among ir galaxies grows along with luminosity (veilleux et al. 1999; goto 2005). in nardini et al. 2010 we have quantified how sf and nuclear activity are the primary engine at the opposite ends of the ulirg luminosity range. the sb component dominates at log(lir/l�) < 12.5, where the agn is a significant contributor in only ∼ 30% of the composite sources. the power supplied by bh accretion grows stronger along the luminosity scale, and ultimately it represents the trigger of the extreme ir activity. ii ) the same technique has been applied by sani et al. 2010 to a sample of local seyfert galaxies to relate their sb activity with the bh growing parameters. we found that sf activity around narrow line seyfert 1 (nls1s), i.e. the more accreting agns, is larger than around bls1 of the same agn luminosity. this result seems to hold over the entire range of distance and luminosity. moreover, the star formation rate is higher in low black hole mass (log(mbh/m�) < 7) and high eddington ratios (l/ledd > 0.1) systems indicating that black hole growth and star formation are occurring simultaneously. iii ) with ir and x-ray analysis we can also investigate the sf–agn connection as a function of obscuration. on one side we do have evidence (veilleux et al. 2009, nardini et al. 2010) that, as expected, the fraction of sb-dominated objects shows a constant decline as the radiation field grows harder with absorption features anti-correlating with the optical classification (siroki et al. 2008). concerning the individual classes, it is worth noting that a sizable number of very powerful agn actually lies among liners (low ionization narrow emitting regions, see also risaliti, imanishi & sani 2010). interestingly seyfert 2, liners and hii regions harbor a similar agn content, supporting the idea of a connection between the sf activity and nuclear obscuration. moreover, sb dominated galaxies host the most obscured agns (see also georgakakis et al. 2004, sani et al. 2008, nardini & risaliti 2011). on the other hand, no enhancement of sf is observed 87 eleonora sani in obscured agns (rovilos et al. 2012) and the specific sf rate of type 2 agn hosts is consistent with the value for normal galaxies (daddi et al. 2007, sarria et al. 2010, mainieri et al. 2011). this is thus a debated matter, but we can not relate the sf physics to the bh physics in a complete and unbiased way because we can not measure e.g. bh mass and accretion rate in type 2 objects. 5 catching feedback in action as mentioned at the beginning of this review, a key ingredient of the merger scenario is feedback. so far a direct detection of the agn winds able to quench sf in the host galaxy is lacking. indeed according to veilleux et al. 2005, ∼ 30% of qso show fast winds (100 km/s < vout < 1000 km/s) but the outflow affects the circumnuclear environment only on the broad/narrow line region scales and thus can not inhibit the sf in the host (muller-sanchez et al. 2011). only in nearby sources like mrk 231 is possible to detect strong molecular outflows (feruglio et al. 2010, fisher et al. 2010) acting on kpc scales as expected by qso-feedback models. the current challenge is to detect such molecular winds at the epoch of the agn activity peak (i.e. at about redshift 2). 6 what we know and what is missing • while the agn accretion history is well traced, we still do not know how the smbh mass function evolves. • if the sf–agn structure on large scales is well studied, we still lack in constrainig the physical processes on small scales (i.e. with extreme adaptive optics corrections for imaging and ifu observations). • we are confident that smbh growth happens simultaneously to the bulk of sf in unobscured composite sources, but the issue for type 2 objects is open. to solve it, we need to identify the entire population and to measure their mbh , l/ledd, sfr etc. • it is well known that agn powers strong 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[51] veilleux, s., kim, d.-c., rupke, d. s. n., et al.: 2009, apj 701, 587. discussion casiana munoz-tunon: more obscured agns have larger sf, and you mention that no correlation is found between sf and obscuration. what sort of connection do you expect? eleonora sani: on the basis of hydrodinamical symulations of major mergers, we do expect the co-presence of a violent sb event and a buried agn at the time of galaxies coalescence. thus a correlation between the sfr and the agn obscuration behaviours (e.g. column density) is expected. analysis based on the cosmos survey don’t find such relation. nonetheless the issue is still a matter of debate since the lack of this correlation could be due to aperture effects. moreover deeply embeded agns are found in local ulirgs and high-redshift smgs. 89 http://dx.doi.org/10.1088/0004-637x/743/1/32 http://dx.doi.org/10.1086/518113 http://dx.doi.org/10.1086/500588 http://dx.doi.org/10.1086/523784 http://dx.doi.org/10.1038/nature08773 http://dx.doi.org/10.1088/0004-637x/768/1/74 introduction evidence of agn–sf connection the major-merger scenario (ulirgs-qso path) the agn vs sf emission x-ray and ir screenings: imaging and spectroscopy the case of ngc 6240 iras 12071-0444: a ct type 2 qso statistical analysis of composite sources catching feedback in action what we know and what is missing 259 acta polytechnica ctu proceedings 1(1): 259–264, 2014 259 doi: 10.14311/app.2014.01.0259 a review of astrophysical jets james h. beall1,2 1st. john’s college, annapolis, md 2space sciences division, naval research laboratory, washington, dc corresponding author: beall@sjc.edu abstract astrophysical jets are ubiquitous: this simple statement has become a commonplace over the last three decades and more as a result of observing campaigns using detectors sensitive from radio to gamma-ray energies. during this epoch, theoretical models of these sources have become more complex, moving from assumptions of isotropy that made analytic calculations possible, to fully anisotropic models of emission from the jets and their interactions with the interstellar and intra-cluster medium. such calculations are only possible because we have extensive computational resources. in addition, the degree of international cooperation required for observing campaigns of these sorts is remarkable, since the instruments include among others the very large array (vla), the very long baseline array (vlba), and entire constellations of satellite instruments, often working in concert. in this paper, i discuss some relevant observations from these efforts and the theoretical interpretations they have occasioned. keywords: astrophysical jets active galactic nuclei uhe cosmic rays quasars microquasars. 1 introduction extended linear structures, usually interpreted as astrophysical jets, are found in star-forming regions, in compact binaries, and of course in active galactic nuclei (agns), and galaxy clusters. gamma-ray bursts can be associated with astrophysical jets associated with stellar collapse where the jets is directed toward earth. we have become aware of considerable details of these results through decades of sustained observing campaigns. the results of these campaigns show similarities and significant differences in the data from some epochs of galactic microquasars, including grs 1915+105, the concurrent radio and x-ray data (beall et al., 1978) on centaurus a (ngc 5128), 3c120 (marscher (2006), and 3c454.3 as reported by bonning et al. (2009), to name a few. the bonning et al. analysis of the 3c454.3 data showed the first results from the fermi space telescope for the concurrent variability at optical, uv, ir, and γ-ray variability of that source. in combination with observations of microquasars and quasars from the mojave collaboration (see, e.g., lister et al., 2009), these observing campaigns provide an understanding of the time-dependent evolution of these sources at milliarcsecond resolutions (i.e., parsec for agns, and astronomical unit scales for microquasars). in blazar sources (see, e.g., ulrich et al. 1997, and marscher, 2006) there seems to be a confirmed connection of jets with accretion disks. in sources without large-scale linear structures (i.e., jets), as ulrich et al. (1997) note, the source variability could result from the complex interactions of the accretion disk with an x-ray emitting corona. but to the extent that ”small” jets are present in these sources, the disk-jet interaction must still be of paramount importance, since it provides a mechanism for carrying away energy from the disk. current theories (see e.g., hawley 2003 and bisnovatyi-kogan 2013 for a discussion of disk structure and jet-launch issues, respectively) suppose that the jet is formed and accelerated in the accretion disk. but even if this is true in all sources, it is still unclear whether or not astrophysical jets with shorter propagation lengths are essentially different in constitution from those that have much longer ranges, or whether the material through with the jet propagates determines the extent of the observational structures we call jets. at all events, the complexities of the jet-ambient medium interaction must have a great deal to do with the ultimate size of emitting region. this sort of argument has applications to both quasars and microquasars, especially if essentially similar physical processes occurs in all these objects (see, e.g., beall, 2003, and marscher, 2006). to some, it has become plausible that essentially the same physics is working over a broad range of temporal, spatial, and luminosity scales. hannikainen (2008) and chaty (2007) have discussed some of the emission characteristics of 259 http://dx.doi.org/10.14311/app.2014.01.0259 james h. beall microquasars, and paredes (2007) has considered the role of microquasars and agns as sources of high energy γ-ray emission. in fact, the early reports of the concurrent radio and x-ray variability of centaurus a can be plausibly interpreted as the launch of a jet from cen a’s central source into the complex structures in its core. additionally, these observations are remarkably similar to the observations of galactic microquasars and agns, including the observations from the fermi space telescope of concurrent γ-ray, ir, optical, and uv variability of 3c454.3 (bonning et al., 2009), and observations by madejski et al. (1997) for bl lac, among others. 2 concurrent variability of agn jet sources as an indication of jet launch or jet-cloud interactions i now turn to two such observations in this paper, the concurrent radio and x-ray variability of centaurus a (beall et al., 1978), and the gamma-ray, uv, and ir concurrent variability discussed by bonning et al., (2009) using the fermi lat and swift spacecraft. 2.1 radio and x-ray variability measurements of centaurus 1 (ngc 5128) the first detection of concurrent, multifrequency variability of an agn came from observations of centaurus a (beall et al., 1978, see figure 1 of that paper). beall et al. conducted the observing campaign of cen a at three different radio frequencies in conjunction with observations from two different x-ray instruments on the oso-8 spacecraft in the 2-6 kev and 100 kev xray ranges. these data were obtained over a period of a few weeks, with the stanford interferometer at 10.7 ghz obtaining the most data. beall et al. also used data from other epochs to construct a decade-long radio and x-ray light curve of the source. figure 1a of beall (2011) shows the radio data during the interval of the oso-8 x-ray observations, as well as the much longer timescale flaring behavior evident in the three different radio frequencies and at both low-energy (2-6 kev, see figure 1b of beall, 2011) and in high-energy (∼ 100 kev, see figure 1c of beall, 2011) x-rays. as noted by beall (2011), a perusal of figure 1a in that paper shows the that the radio data (represented as a “+” in that figure) generally rise during 1973 to reach a peak in mid-1974, then decline to a relative minimum in mid-1975, only to go through a second peak toward the end of 1975, and a subsequent decline toward the end of 1976. this pattern of behavior is also shown in the ∼ 30 ghz data and the ∼ 90 ghz data albeit with less coverage at the higher two radio frequencies. several points are worthy of note. first, as beall et al. (1978) show, the radio and x-ray light curves track one another. this result demonstrated the first report of concurrent radio and x-ray variability of an active galaxy. mushotzky et al. (1978), using the weekly 10.7 ghz data obtained by beall et al., (1978) demonstrate that the 10.7 ghz radio data track the 2-6 kev x-ray data on weekly time scales, also. the concurrent variability at radio and x-ray frequencies argues that the emitting region is the same for both the radio and xray light. this, as was noted by beall and rose (1980), can be used to set interesting limits on the parameters of the emitting region. in addition, the observations at the three radio frequencies (10.7 ghz, ∼ 30 ghz, and ∼ 90 ghz) clearly track one another throughout the interval whenever concurrent data are available. figure 1: fermi lat data (see, e.g., lott et al. 2013) showing the time history of the flaring from 3c454.3 (figure 1b), along with the bonning et al. (2009) data showing concurrent gamma-ray, ir, optical, and uv data (figure 1b). the time of the jet launch from the mojave ”movie” is illustrated by the arrow. the result, first reported here, shows radio flaring consistent with the bonning et al. data. a plausible hypothesis for the observations we have witnessed is that they arise from physical processes associated with the ”launch” of an astrophysical jet into the complex structures in the core of centaurus a. the timescale of the evolution from early 1973 through 1977 appears to be associated with the evolution of a larger structure over a more extended region. the observations are consistent with the interaction of the astrophysical jet with an interstellar cloud in the core of cen a. 260 a review of astrophysical jets it is clear from this discussion that a distinction needs to be made about which observational signatures are associated with the jet launch, the jet itself, and the ambient medium’s reaction to the jet. in considering such a scenario applied to microquasars, the observations of sco x-1 by fomalont, geldzahler, and bradshaw (2001), as discussed by beall et al. (2013) are extremely informative. 2.2 a jet launch coincident with the 3c454.3 multi-frequency flaring bonning et al. (2009) performed an analysis of the multi-wavelength data from the blazar, 3c454.3, using ir and optical observations from the smarts telescopes, optical, uv and x-ray data from the swift satellite, and public-release γ-ray data from the fermilat experiment. in that work, she demonstrated an excellent correlation between the ir, optical, uv and gamma-ray light curves, with a time lag of less than one day. urry (2011) noted that 3c454.3 can be a laboratory for multifrequency variability in blazars. while a more precise analysis of the data will be required to determine the characteristics of the emitting regions for the observed concurrent flaring at the different frequencies, the pattern of a correlation between low-energy and higher-energy variability is consistent with that observed for cen a, albeit with the proviso that the energetics of the radiating particles in 3c454.3 is considerably greater. a perusal of the data for 3c454.3 at milliarcsecond scales taken from the mojave vlba campaign during the bonning et al (2009) flare show that the timedependent flare occurs during launch of a new component of the jet that originates from the core. figure 2: figure 2: 3c454.3 shown at milliarcsecond scales for data taken from the mojave vlba campaign during the bonning et al (2009) flare. a perusal of the time-dependent flare clearly shows that the jet launch from the core of the agn is coincident with the double peaked flaring shown in the bonning et al. results. the pattern of variability reported by bonning et al., is consistent with the injection of relativistic particles into a region with relatively high particle and radiation densities (i.e., an interstellar cloud). the picture that emerges, therefore, is consistent with the observations of spatially and temporally resolved galactic microquasars and agn jets. it should not escape our notice that other epochs in the 3c454.3 data from mojave and fermi also show similar periods of injection of radio blobs that are associated with a double-peaked structure in the gamma-ray light. for the cen a data, and for data from 3c454.3, it is the concurrent variability that suggests that the radio to x-ray (in cen a’s case) and the ir, optical, and uv to γ-ray fluxes (in 3c454.3’s case) are created in the same region. this leads to the possibility of estimates of the source parameters that are obtained from models of these sources. vlbi observations of cores vs. jets (see, e.g., the study of bl lac by bach et al. 2006) show the structures of the core vs. jet as they change in frequency and time. it has thus become possible to separate and study the time variability of the jet and the core of agn at remarkably fine temporal and spatial scales. van der laan (1976) discussed the theoretical interpretation of cosmic radio data by assuming a source which contained uniform magnetic field, suffused with an isotropic distribution of relativistic electrons. the source, as it expanded, caused an evolution of the radio light curve at different frequencies. each of the curves in van der laan’s paper represents a factor of 2 difference in frequency, the vertical axis representing intensity of the radio flux and the horizontal axis representing an expansion timescale for the emitting region. van der laan’s calculations show a marked difference between the peaks at various frequencies. the data from cen a (as discussed more fully in beall (2008, 2010) are, therefore, not consistent with van der laan expansion (van der laan, 1976), since for van der laan expansion, we would expect the different frequencies to achieve their maxima at different times. also, the peak intensities should decline with increasing frequencies at least in the power-law portion of the spectrum. the most likely explanation for the changes in the spectrum of the cen a data at 100 kev (beall, et al., 1978) is that the emitting region suffered an injection of energetic electrons. that is, a jet-ambient-medium interaction dumped energetic particles into a putative ”blob,” or, equivalently, that there was a re-acceleration of the emitting electrons on a timescale short compared to the expansion time of the source. an analysis of the 3c120 results compared with the data from the galactic microquasar, sco x-1, under261 james h. beall taken by beall (2006) shows a similar radio evolution, with rapidly moving “bullets” interacting with slower moving, expanding blobs. it is highly likely that the elements of these sources that are consistent with van der laan expansion are the slower-moving, expanding blobs. i believe that the relativistically moving bullets, when they interact with these slower-moving blobs, are the genesis of the flaring that we see that seems like a re-acceleration of the emitting, relativistic particles. i note that a similar scenario could be operating in cen a and 3c454.3. this is not to say that the ”slower-moving blobs” are not themselves moving relativistically, since the bipolar lobes have significant enhancements in brightness due to relativistic doppler boosting for the blobs moving toward us. the true test of this hypothesis will require concurrent, multifrequency observations with resolutions sufficient to distinguish jet components from core emissions in galactic microquasars as well as for agn jets. one of the most remarkable sagas regarding the discovery of quasar-like activity in galactic sources comes from the decades long-investigation of sco x-1 by ed fomalont, barry geldzahler, and charlie bradshaw (fomalont, geldzahler, and bradshaw, et al. 2001). during their observations, an extended source changed relative position with respect to the primary object, disappeared, and then reappeared many times. we now know that they were observing a highly variable jet from a binary, neutron star system. the determinant observation was conducted using the very large array (vla) in socorro, new mexico and the vlba interferometer (see, e.g., beall, 2008) for a more complete discussion). the observations of the concurrent ir, optical, uv, and γ-ray variability of 3c454.3, and its associated jet launch in the mojave data, argue for a reinvestigation of these data sets in the near future. it is worthy of note that the milliarcsecond observations show a complex evolution of structure at parsec scales, including an apparently sharp change of directions associated with changes in the polarization of the radio light at that point in the jet’s evolution (see figure 2). furthermore, the multi-frequency flares reported by bonning et al. (2009) are consistent with the launch of another component of the astrophysical jet in the core region. regarding the acceleration region and the possible mechanisms for the collimation of the jets, a number of models have been proposed (see, e.g., kundt and gopal-krishna (2004), bisnovaty-kogan et al. (2002), romanova and lovelace (2009), and bisnovaty-kogan (2012) that might help explain the complexity present in these data. 3 concluding remarks the data discussed therein suggest a model for the jet structures in which beams or blobs of energetic plasmas propagate outward from the central engine to interact with the ambient medium in the source region. this ambient medium in many cases comes from prior ejecta from the central source, but can also come from clouds in the broad line region. the jet can apparently also excavate large regions, as is suggested by the complex structures in, for example, 3c120. the physical processes which can accelerate and entrain the ambient medium through which the jet propagates, have been discussed in detail in several venues (see, e.g., rose et al., 1984, 1987, beall, 1990, beall et al., 2003, and beall, 2010). in this paper, we have focused on the patterns of concurrent radio and x-ray variability for cen a, and on concurrent radio, gamma-ray, optical, ir, and uv variability for 3c454.3. the data can be interpreted as being associated with a jet-launch scenario for these sources, and this paper represents the first report of the association between a jet launch in the mojave data for 3c454.3 and its gamma-ray flare from fermi. in addition, jets from microquasars show similar patterns of variability to those of agns. these data require us to abandon our assumptions both of spherical symmetry and of single-zone productions in our models of these sources. acknowledgement the author gratefully acknowledges the support of the office of naval research for this research. this research has made use of data from the mojave database that is maintained by the mojave team (lister et al., 2009, aj, 137, 3718). the author gratefully acknowledges the data provided by the mojave team for this paper. references [1] the pierre auger collaboration, et al., 2007, science 318, 938. doi:10.1126/science.1151124 [2] bach, u., villata, m., raiteri, c. m., et al. 2006, ”structure a variability in the vlbi jet of bl lacertae during the webt campaigns (1995-2004),” a&a, 456, 105. 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[46] zanni, c., murante, g., bodo, g., massaglia, s., rossi, p., ferrari, a., 2005 astron. astrophys., 429, 399. discussion todor stanev do you expect an important role of magnetic reconnection in the particle acceleration in the jets? jim beall i do believe that magnetic reconnection plays an important part in particle acceleration, not only in the launch region, but also in the jet itself. if you recall the data i showed for 3c454.3, the region where the get seems to turn shows marked variations in brightness that seem to be associated both with the blobs from the jet traversing that region, and with the jet luminosity at the ”elbow.” in fact, it seems likely that the acceleration and collimation region of the jet in 3c454.3 extends many parsecs from the launch region. carlotta pittori can you say something about the southern jet of the crab nebula? as i actually showed at a previous vulcano workshop and promptly noticed by w. kundt, the jet orientation for x-ray images from chandra changed in the ten years from 20002010. how can this be explained? by kink instabilities? where does the field close in this system? jim beall this is only a conjecture, but if we look at the large distance where the jet in 3c454.3 changes direction, it appears that magnetic field structures can extend out to several parsecs and maintain their coherence, at least in agn sources. if the same is true in the crab, then the magnetic fields might be responsible for the changes in the jet orientation. wolfgang kundt can you tell me when and where geoff burbidge argued in favor of hadronic jets on the grounds of their stability? in my model with gopal, a pair-plasma jet is thought to be stable on the grounds of their exb drift, for example in new astronomy reviews, 2008, 52, pp 364-369. jim beall geoff burbidge’s article was published in astrophysical journal, 1956, 124, 416. that paper was in the main an argument that the energy supplied to the giant radio lobes must be of order 1060 ergs, and could only be delivered by hadronic jets. 264 introduction concurrent variability of agn jet sources as an indication of jet launch or jet-cloud interactions radio and x-ray variability measurements of centaurus 1 (ngc 5128) a jet launch coincident with the 3c454.3 multi-frequency flaring concluding remarks 13 acta polytechnica ctu proceedings 1(1): 13–19, 2014 13 doi: 10.14311/app.2014.01.0013 γ-rays, ν’s and particle astronomy as messengers of the universe maurizio spurio1 1dipartimento di fisica e astronomia università di bologna and infn corresponding author: spurio@bo.infn.it abstract observations of charged cosmic-rays, γ-rays and neutrino are possible thanks to the availability of new detectors coming from technologies typical of experimental particle physics. in conjunction with more traditional techniques used by astronomers, these multimessenger correlations of experimental data are opening a new scenario in astrophysics. we review some of the more recent developments in the field. keywords: cosmic-rays γ-rays neutrinos satellites ground-based experiments neutrino telescopes. 1 introduction the multimessenger inter-correlation between cosmic rays (crs), γ-rays and neutrinos is of fundamental importance for a deeper comprehension of high energy (he) processes in astrophysical sources. one of the main questions in astrophysics is the origin and nature of crs (§2), whose energy spectrum extends up to above 1020 ev. there are many indications of the galactic origin of the cr bulk (protons and other nuclei up to ∼ 1016 ev). due to the influence of galactic magnetic fields, charged particles detected on earth do not point to the sources. gamma-rays and neutrinos do not suffer the effect of magnetic fields: they represent the decay products of accelerated charged particles but cannot be directly accelerated. the possibility of particle astronomy rely on the detection of these neutral probes, or to the detection of the highest component of charged crs (see castellina, these proceedings). recent advances on gev and tev γ-ray astronomy by satellites and ground-based experiments (§3) have led to the discovery of different classes of galactic and extragalactic objects that can be shed light on the cr acceleration mechanisms. most objects observed in the γ-rays (see fig. 1) are known sources from measurements of their electromagnetic emission at different wavelengths, in particular in the radio and in the x-rays. accelerated protons will interact in the surroundings of the crs emitter with photons predominantly via the ∆+ resonance: p + γ → ∆+ → πo + p or → π+ + n (1) protons will interact also with ambient matter (protons, neutrons and nuclei), giving rise to the production of charged and neutral mesons. the relationship between sources of γ-ray and neutrinos is the meson-decay channel. neutral mesons decay in πo → γγ while charged mesons decay in neutrinos: π+ → νµ + µ+ ↪→ µ+ → νµ + νe + e+ π− → νµ + µ− ↪→ µ− → νµ + νe + e− (2) in particular circumstances, the energy escaping from the source is distributed between crs, γ-rays and neutrinos. only the coincident measurement of neutrinos (§4) from the source would give a uncontroversial proof of the discovery of the galactic acceleration sites of protons and nuclei. figure 1: sketch of the production of crs, neutrinos, γ-rays and lower energy photons from an astrophysical jet. the highest energy crs are probably originated from extragalactic sources. if they are protons with e > 5 × 1019 ev, they could interact with the cosmic microwave background radiation. this process restricts the origin of highest energy protons seen on earth to a small fraction of the universe, a sphere of the order of 13 http://dx.doi.org/10.14311/app.2014.01.0013 maurizio spurio 100 mpc (the greisen-zatsepin-kuzmin suppression). the prediction of a diffuse flux of he neutrino of extra-galactic origin is a direct consequence of the ultra he cr observations. this connection between crs, neutrinos and γ-rays was used to put upper bounds on the expected neutrino flux from extragalactic sources, since the neutrino energy generation rate will never exceed the generation rate of high energy protons. in this paper, we will review some of the recent updates in these fields. 2 crs and antimatter figure 2: cr intensity as a function of the energy from 1010 to 1020 ev. below few gev, the contribution of particles coming from the sun is not negligible. the energy range of the flux measured by direct experiments is reported as blue line; that measured by indirect experiments as red line crs are mainly protons (fig. 2) and heavier nuclei which are constantly hitting the upper shells of the earth’s atmosphere (blümer et al, 2009). up to energies of ∼ 1014 ev, the cr spectrum is directly measured above the atmosphere. stratospheric balloons and satellites have provided the most relevant information about the composition of crs. above ∼ 1014 ev, cr measurements are only accessible from groundbased large area arrays. recent observations of tev photons seem to confirm the diffusive shock acceleration (dsa) model, where the shocks powered by supernova explosions in our galaxy provide the iterative scattering mechanism for crs acceleration up to ∼ 1016 ev (see hillas, 2005). the key feature of this process is that an energy powerlaw spectrum of the type ∼ e−2 is produced. the dsa model is consistent with the balance between the energy transferred to the accelerated particles and the energy loss due to the escape of crs out of the galaxy. the e−2 emission at sources, when corrected for energydependent leakage from the galaxy, is in agreement with the observed cr energy spectrum. the details of acceleration mechanism and propagation of cosmic rays at higher energies are not completely understood. an important open item is the presence of antimatter in the radiation. equal amounts of matter and antimatter should have been produced at the beginning of the universe. the fact that there seems to be only matter around us is one of the major open problems in cosmology and in particle physics. the antiprotons, as well as the positrons, are a component of the cosmic radiation being produced in the interaction between crs and the interstellar matter. being exactly the same as particles except for their opposite charge sign, antiparticles are readily distinguished as they bend in opposite directions in magnetic field. magnetic spectrometers provide a clear and simple particle/antiparticle separation and probe the existence of antimatter in our galaxy. figure 3: the positron ratio measured by the pamela, fermi and ams-02 as presented in the first public presentation of ams-02 results. the expectation from the production of positrons as secondary particles is shown as dashed line. the discrepancy below 10 gev is explained by different phases of solar activity when data were collected. from http://ams.cern.ch/ an important feature in the e+ spectrum was observed by the pamela satellite and confirmed with the high precision measurement of ams-02. launched in 2011, the ams-02 experiment started immediately to take date on the international space station (iss). it is the largest particle physics detector never carried outside the earth atmosphere. the first results were presented in april, 2013 and refer to the measurement of positron in the crs. the e+ fraction in the e− + e+ flux found in the cosmic radiation increases steadily be14 γ-rays, ν’s and particle astronomy as messengers of the universe tween 10 gev to ∼ 250 gev, see fig. 3. at energies above 250 gev, the spectrum appears to flatten but to study the behavior at higher energy more statistics is required. the positron fraction spectrum exhibits no spatial anisotropy, structures or time dependence. these measurements have stimulated a large debate, as they cannot in fact be understood by models describing the production of secondary crs during propagation in the galaxy (see dashed line in fig. 3). several explanations have been proposed to interpret the observed excess: an astrophysical origin, such as nearby pulsars or microquasars, or exotic sources, as for instance the annihilation of dark matter particles in our galaxy. the p/p ratio (∼ 10−5 ÷ 10−4) shows that the antiproton flux is in overall agreement with a pure secondary component: antiparticles are produced during the interaction of crs with the interstellar matter. on the contrary, the ratio e+/e− ∼ 0.1 seems indicate that most of the detected electrons are of primary origin. due to energy loss processes, the majority of he electrons must be originated by sources closer than a few hundred pc. he electrons really probe cr production and propagation in the nearby region of our galaxy. in the next years, the ams-02 experiment will undoubtedly be the leading experiment for a systematic study of the crs through direct measurements, for the searches of antimatter in space, and for the searches for particles originated by dark matter annihilations. 3 high energy γ-rays he astrophysical processes producing relativistic particles are in most cases associated with the production of γ-rays in a wide range of energies. starting from the egret satellites in the ’90th, a new window has been opened in the observation of photons above ∼ 100 mev 1. the egret catalog consists of about 270 galactic and extragalactic objects. the present generation of space-based gamma ray telescopes, agile and lat, has opened a new era with thousands of galactic and extragalactic sources. the large number of sources and the high data quality are producing a deeper insight into the understanding of the processes of acceleration and radiation of non-thermal particles in the universe at sub-tev energies. as the flux of γ-rays decreases with increasing energy, space-based observations of most sources are limited to less than 100 gev. gamma-rays above 100 gev 2 are detected on ground, using extensive air shower (eas) particle detectors or the imaging atmospheric-cherenkov technique (iact). tev γ-rays are absorbed when reaching the earth atmosphere, and the absorption process proceeds by creation of a cascade of secondary particles. these emit cherenkov radiation, at a characteristic angle in the visible and uv range, which passes through the atmosphere. as a result of cherenkov light collection by a suitable mirror in a camera, the showers can be observed on the surface of the earth. the comparisons between the main features of space-based, iacts, eas and neutrino telescopes (§4) are presented in table 1. table 1: characteristic of space-based, iacts, eas and neutrino telescopes. from the top raw: energy range; effective area; the background rejection power; the angular and energy resolutions; the characteristic aperture of the telescopes; the duty cycle. space iact eas ν-tel energy (tev) 10−4-0.3 0.5-100 0.5-50 1-1000 area (m2) 1 104 104 0.1-300 bck rejec. > 99% > 99% 95% 99% angular resol. 0.5◦ 0.05◦ 0.7◦ 0.5◦ ∆e/e 10% 15% 50% 40% aperture (sr) 2.7 0.003 1.8 2π duty cycle 85% 10% 95% 95% 3.1 he γ-rays on satellites the large area telescope (lat) on the fermi satellite (launched in june 2008) is a γ-ray detector designed to distinguish photons in the energy range 20 mev to more than 300 gev from the high background of energetic charged particles. the γ-ray sky (unlike at other wavelengths) is strongly dominated by diffuse radiation originating in our galaxy by cr interactions with the interstellar gas and photon fields through the processes of inelastic nucleon scattering, bremsstrahlung, and inverse compton scattering. the lat mapping of the galactic γ-rays offers a way to derive information about the spatial distribution of crs and matter. the second catalog of high-energy γ-ray sources (2fgl) detected by the lat (nolan p.l. et al., 2012) derives from data taken during the first 24 months of the science phase of the mission, which began on 2008 august 4. the 2fgl catalog contains 1873 sources detected and characterized in the 100 mev to 100 gev. it includes source location regions, energy dependence of the flux in terms of a power-law with an exponentially cutoff and light curves on monthly intervals for each source. among the 1873 sources, 127 are firmly identified and 1171 reliably associated with counter1we indicate with he γ-rays the photons between 0.02-100 gev 2we indicate with vhe γ-rays the photons between 100 gev-100 tev 15 maurizio spurio parts of known or likely γ-ray producing source classes, see fig. 4. agn, and in particular blazars, are the most prominent class of associated sources: 917 sources are associated, of which 894 are blazars, 9 are radio galaxies, 5 seyfert galaxies. normal galaxies are now established as a class of γ-ray emitters and 7 2fgl sources are associated with such objects. pulsars have been traditionally studied through radio astronomy methods, with about 1800 pulsars found beaming radio waves. however, most of their radiation (a few percent of their spin-down power) is emitted at high-energies. in the last few years, the number of pulsars detected in the γ-ray has increased from half a dozen to more than 150 thanks to the agile satellite and the fermi-lat. among the pulsar, the crab plays a fundamental role, as it is the strongest γ-ray source in our field of view. the crab was since recently thought to be a steady source of radiation, from optical to tev energies. agile unexpected has discovered in 2010 (tavani et al., 2011) strong and rapid γ-ray flares from the crab nebula over daily timescales. this observation have changed the understanding of this cosmic object, and challenged emission models of pulsar wind and particle acceleration processes. pulsed γ-ray emission above 100 gev and up to 400 gev was recently detected from the crab with the veritas and magic iacts. figure 4: fermi-lat 2fgl full sky map (top) and blowup of the inner galactic region (bottom) showing sources by source class. identified sources are shown with a red symbol, associated sources in blue. about 70 sources of the 2fgl catalog are associated with pulsar wind nebulae. supernova remnants are a special class because a substantial number of the known objects are sufficiently extended to be potentially resolved with the lat. these objects are particularly interesting because they can represent regions of cr acceleration. in total, twelve sources in the catalog are modeled as spatially extended. 3.2 vhe γ-rays at ground one of the most recent and remarkable achievements in astrophysics is the discovery of more than 110 galactic and extragalactic sources of vhe radiation (see for a review rieger et al., 2013). at present, seven shell-type snrs have been firmly identify as vhe γ-ray emitters the pioneering ground based γ-ray experiment was built by the whipple collaboration. during the last decade, several ground-based γ-ray detectors were developed, both in the north and south earth hemisphere. at present, the new generation apparatus are the h.e.s.s. and veritas telescope arrays and the magic telescopes. these iacts have produced a catalogue of tev γ-ray sources which is continuously updated and available at http://tevcat.uchicago.edu/. electron can produce he γ-rays through the socalled leptonic model. synchrotron radiation from radio to the x-ray band is originated by accelerated electrons moving in the source magnetic fields. these particles can also produce gev-tev γ-rays through inverse compton scattering on the produced radiation field, or on external radiation fields. therefore, measurements of the synchrotron x-ray flux from a source can constrain the predictions on the accompanying γ-rays produced in leptonic processes. accelerated hadrons could either produce γ-rays via interaction with ambient matter or photon fields with sufficient high density. in this case, γ-rays are produced by the decay of neutral mesons while ν’s are produced by the decay of charged mesons. this γ-rays and neutrino production refer to a so-called astrophysical hadronic model. in this framework, the energy spectrum of secondary γ and ν particles follows the same power law of the progenitor crs. both the leptonic and the hadronic models, or a combination of them, could provide an adequate description of the present experimental situation. if high energy photons are produced in the hadronic models, high energy neutrinos will be produced as well. most of observed tev γ-ray galactic sources have a power law energy spectrum e−αγ , where αγ ∼ 2.0÷2.5. the values of the spectral index are very close to the expected spectral index of cr sources, αcr. this lead to the conclusion that sources of tev γ-rays can also be the sources of galactic crs. 16 γ-rays, ν’s and particle astronomy as messengers of the universe 4 neutrino astrophysics the main detection signatures for neutrinos in a neutrino telescope (see chiarusi & spurio, 2010) are long, straight tracks and approximately spherical cascades. the former are created by neutrino-induced muons while the latter are produced by neutrino-induced electromagnetic and/or hadronic showers. charged particles emit cherenkov light. from the measured arrival time of the cherenkov light, the direction of the neutrino can be derived. the accurate measurement of the νµ direction (up to 0.3 o in water) could allows the association with (known) sources. because the mechanisms that produce crs can produce also neutrinos and γray, potential neutrino sources are in general also γ-ray emitters. due to the production mechanism (eq. 2), the flavor ratio at sources is νe : νµ : ντ = 1 : 2 : 0, which is changed by the neutrino oscillations to 1 : 1 : 1 on earth. figure 5: neutrino telescopes searches for an excess of events over the irreducible background of atmospheric neutrinos. the points represent the measurements of atmospheric νµ spectrum by three telescopes. the red curve represent the still unmeasured contribution of prompt neutrinos from charmed mesons decay. the green lines are the waxman& bahcall (1999) upper bound from diffuse flux of neutrinos from extragalactic sources. the blue line the possible contribution of ν from the gzk effect. some of the most promising candidate neutrino sources in our galaxy are extremely interesting, due to the recent results from tev γ-ray detectors. a neutrino telescope in the mediterranean sea is looking at the same southern field-of-view as the h.e.s.s. including the galactic center. the small interaction cross section of neutrinos allows them to come from far away, but it is also a drawback, as their detection requires a large target mass. assuming the present boundaries arising from γ-rays sources, the challenge to detect galactic neutrinos is open for a multi kilometer-scale apparatus, see fig. 5. also the extragalactic crs-neutrinos connection sets the scale of the detectors to 1 km3. icecube (http://icecube.wisc.edu/) is a 1 km3-scale neutrino detector buried in the antarctic ice. it comprise 86 strings, with 5160 photomultiplier tubes (pmt). each string includes 60 digital optical modules. the deepcore infill array to icecube reduces the energy threshold of icecube to energies as low as 10 gev. in water the antares collaboration (ageron et al, 2011) has completed in 2008 the construction of the largest neutrino telescope (∼ 0.1 km2) in the northern hemisphere. galactic sources. although much smaller than icecube, antares is advantaged by its geographical location for the study of galactic sources. to give a figure of merit, the tev photon flux from a possible neutrino candidate source as rx j1713.7-3946 is e2γφγ ' 10−11 tev cm−2 s−1. assuming the same flux from neutrinos and no he cutoff, few events/year are expected in a 1 km3 telescope over the background of atmospheric neutrinos in a ∼ 1◦ search cone. the sensitivity of antares for 1000 days of livetime is ∼ e2νµ φ sens νµ ' (3 − 8) × 10−11 tev cm−2 s−1 in the declination range from δ = −90◦ to +43◦. in total 51 possible neutrino sources have been studied. none of them shows a significantly excess of tev events over the expected background. figure 6: expected neutrino spectra from full numerical neucosma (red) and simply analytic guetta (blue) models. limits on these predictions are shown in the energy ranges where 90% of the flux (dashed lines) is expected. the icecube limit on the neutrino emission (black dashed) is based from 300 grbs. the antares limit on 296 grbs (red dashed) takes into account the neutrino oscillation effect. ν from grbs. gamma-ray bursts (grbs) are short and very intense flashes of he γ-rays, which occur unpredictably and isotropically over the sky. in model describing the γ-ray emission, protons can also be shockaccelerated, yielding secondary emission of he ν’s accompanying the electromagnetic signal. the detection of neutrinos in coincidence with a grb would be unambiguous proof for hadronic acceleration in cosmic sources. 17 maurizio spurio fig. 6 shows the result from these searches. the stringent icecube limit (black dashed line, abbasi et al., 2012) was a factor of 3.7 below the predictions made using the guetta et al. (2004) model (blue line in fig. 6), creating tension between the nonobservation of a signal and the prevailing models for neutrino emission from grbs. this could either indicate the need for rejection of these models or for more detailed modeling of the neutrino emission. the ν prediction from an advanced numerical calculation of grb (neucosmahümmer et al., 2012), including the full photohadronic interaction cross-section, independent losses of secondary particles, and flavor mixing, is shown as read line. the new predicted flux is about an order of magnitude below than the previous analytic approach, and thus it is still compatible with the limit as published by icecube. figure 7: declination vs reconstructed energy of the he cosmic neutrino candidates in the icecube detector (whitehorn at wipac 2013) diffuse he neutrinos. recently icecube reported on the observation of two he particle shower events discovered in a search for pev neutrinos (aartsen et al., 2013). these events (reconstructed as downward going) are of much higher energy than expected from the background of atmospheric neutrinos, fig. 5. stimulated by these two events, almost during the same time of this frascati workshop, the icecube announced the results of a dedicated search for downward going neutrino-induced events in a restricted fiducial volume. here, 28 events (7 muon-like and 21 shower-like) were discovered, with an expected background of 12.1 events (from atmospheric and prompt neutrinos, and atmospheric muons). the distribution of the declination and deposited energy for these events is shown in fig. 7. the collaboration asserts that the events seems to be neutrinos, with flavor ratios consistent with the expected 1 : 1 : 1 and compatible with an isotropic flux. 5 conclusions multimessenger astrophysics becomes more and more important to obtain a complete picture of non-thermal processes in the universe. this program can be achieved by combining pieces of information from all three messengers: photons, charged crs and ν’s. the implications extend from the origin of crs to the origin of dark matter, from processes of acceleration of particles by strong shock waves to the magnetohydrodynamics of relativistic jets, from distribution of matter in the interstellar medium to the intergalactic radiation and magnetic field distributions. the strong impact of he and vhe γ-rays discoveries on several topical areas of modern astrophysics and cosmology are recognized and of fundamental importance for the astronomical communities. in the near future, the role of ams-02 is of overwhelming importance for direct measurement of crs. finally, the recent multi-tev excess of neutrino events may be a first hint of an astrophysical he neutrino flux, opening the field of neutrino astronomy for the next decade. references [1] aartsen m.g. et al., 2013. accepted by prl [2] abbasi, r. et al.: 2012, nat, 484, 351 doi:10.1038/nature11068 [3] ageron, m. et al.: 2011, nim a, 656, 11 doi:10.1016/j.nima.2011.06.103 [4] blümer j. et al: 2009, prog.part.nucl.p. 63, 293 doi:10.1016/j.ppnp.2009.05.002 [5] chiarusi t, spurio m.: 2010, epj. c65, 649 [6] guetta, d. et al.: 2004, astr. phys., 20, 429 doi:10.1016/s0927-6505(03)00211-1 [7] hillas a. m.: 2005, j. phys. g 31, 39. doi:10.1088/0954-3899/31/5/r02 [8] hümmer, s. et al.: 2012, prl 108, 231101 doi:10.1103/physrevlett.108.231101 [9] nolan p.l. et al.: 2012 apjs 199 31 doi:10.1088/0067-0049/199/2/31 [10] rieger f.m. et al.:arxiv:1302.5603 [11] tavani m. et al.: 2011, science vol. 331 doi:10.1126/science.1200083 [12] waxman e., bahcall j.: 1999, prd59:023002 18 http://dx.doi.org/10.1038/nature11068 http://dx.doi.org/10.1016/j.nima.2011.06.103 http://dx.doi.org/10.1016/j.ppnp.2009.05.002 http://dx.doi.org/10.1016/s0927-6505(03)00211-1 http://dx.doi.org/10.1088/0954-3899/31/5/r02 http://dx.doi.org/10.1103/physrevlett.108.231101 http://dx.doi.org/10.1088/0067-0049/199/2/31 http://dx.doi.org/10.1126/science.1200083 γ-rays, ν’s and particle astronomy as messengers of the universe discussion arnon dar: concerning the icecube events, the quoted error on the atmospheric neutrino background seems to be very small. maurizio spurio: in my opinion the main uncertainty could arise not from atmospheric neutrinos, but from the vetoing efficiency estimate for surviving atmospheric muons. there are some peculiarity of the events which are intriguing and only a detailed publication can clarify. for instance, the fact that (after background subtraction) the muon neutrinos are fewer than the expected from the quoted 1 : 1 : 1 ratio; that a cutoff energy at ∼1 pev should be present; that the upgoing signal (4 events) is compatible with the background (3 events). for these reasons, if there is no underestimated contamination from atmospheric muons, the signal seems to be more likely of galactic origin than diffuse extragalactic. 19 introduction crs and antimatter high energy -rays he -rays on satellites vhe -rays at ground neutrino astrophysics conclusions 175 acta polytechnica ctu proceedings 1(1): 175–180, 2014 175 doi: 10.14311/app.2014.01.0175 be/x-ray binaries with black holes in the galaxy and in the magellanic clouds janusz zió lkowski1 1copernicus astronomical center, ul. bartycka 18, 00-716 warsaw, poland corresponding author: jz@camk.edu.pl abstract i will start with the statistics indicating that the objects named in the title of my talk are either non-existing or very elusive to detect (not a single such object is known against 119 known be/neutron star x-ray binaries). after brief reviewing of the properties of be/x-ray binaries i discuss several objects that were proposed as the long sought for candidates for be/black hole x-ray binaries. after three unsuccessful candidates (ls i +610 303, ls 5039 and maxi j1836-194), a successful candidate (agl j2241+4454/mwc 656) was finally, very recently, announced. keywords: stars: binaries – stars: x-ray binaries – stars: be/x-ray binaries – stars: black holes – stars: be. 1 introduction the objects named in the title of my talk seem not to exist neither in our galaxy nor in the magellanic clouds. not a single such object is known today (this statement was still true during our conference). at the same time, we know (today) 184 be/x-ray binaries (74 in the galaxy and 110 in the magellanic clouds). 2 properties of be/x-ray binaries be x-ray binaries (be xrbs) are the most numerous subclass among high mass x-ray binaries. they outnumber all other high mass x-ray binaries by a factor of three (we know about 180 be xrbs vs about 60 other high mass x-ray binaries). these systems consist of a be star and a compact object (a neutron star or a black hole). the be stars are massive, generally main sequence, stars of spectral types a0-o8 with balmer emission lines (negueruela 1998). the be xrbs are rather wide systems (orbital periods in the range of ∼ 10 − 1180 days). the orbits are frequently eccentric. a compact component accretes from the wind of a be star (even massive be stars are well within their roche lobes for these wide orbits, so there is no question of roche lobe overflow). at present, 184 be xrbs are known in the galaxy and in the magellanic clouds, and in 119 of them, the compact object was confirmed to be a neutron star by the detection of the x-ray pulsations (with the pulse periods in the range of 34 ms to ∼ 1400 s). in the remaining cases, whenever we have information concerning the nature of the compact component (such as an x-ray spectrum), it also indicates a neutron star. although one cannot exclude that a few of these systems contain white dwarfs or black holes, it is safe to state that majority of them contain a neutron star as compact component. however, not a single black hole binary containing a be type component has been found so far. this disparity (119 be xrbs with neutron stars versus not a single one with a black hole) seems indeed striking (this was the situation during our conference; today the statistics is 119 to one − see the section 4.2). the x-ray emission from be xrbs (with a few exceptions) is of a distinctly transient nature with rather short (days to weeks) active phases separated by much longer (months to tens of years) quiescent intervals (a typical flaring behavior). there are two types of flares, which are classified as type i outbursts (smaller and roughly regularly repeating) and type ii outbursts (larger and irregular. this classification was defined by negueruela & okazaki 2001 and negueruela et al. 2001. type i bursts are observed in systems with highly eccentric orbits. they occur close to periastron passages of a neutron star. they are repeating at intervals ∼ porb. type ii bursts may occur at any orbital phase. they are correlated with the disruption of the excretion disc around be star (as observed in hα line). they repeat on time scale of the dynamical evolution of the excretion disc (∼ few years to few tens of years). this recurrence time scale is generally much longer than the orbital period (negueruela et al. 2001). be xrbs systems are known to contain two discs: excretion disc around be star and accretion disc around neutron star. both discs are temporary: excretion disc disperses and refills on time scales ∼ few years to few decades (dynamical evolution of the be star disc, formerly described as ”the activity of a be star” (negueru175 http://dx.doi.org/10.14311/app.2014.01.0175 janusz zió lkowski ela et al. 2001)), while the accretion disc disperses and refills on time scales ∼ weeks to months (which is related to the orbital motion of a neutron star on an eccentric orbit and, on some occasions, also to the major instabilities of the other disc). the accretion disc might be absent over a longer period of time (∼ years), if the other disc is very weak or absent. the x-ray emission of be xrbs binaries is controlled by the centrifugal gate mechanism, which, in turn, is operated both by the periastron passages (type i bursts) and by the dynamical evolution of the excretion disc (both types of bursts). this mechanism explains the transient nature of the x-ray emission ( see zió lkowski 2002 and references therein). the more detailed description of the properties of be xrbs is given, e.g. in negueruela et al. 2001, zió lkowski 2002, belczyński & zió lkowski 2009, reig 2011 and references therein. considering the numbers of different (neutron star vs black hole) be xrbs, it is important to note, that if we take all other xrbs (both high and low mass), except be xrbs, then the ratio of neutron star systems to black hole systems is about 2:1 (remillard & mcclintock 2006, zió lkowski 2008). for be xrbs this ratio is 119 to zero. the disparity is indeed striking. this disparity is referred to as a missing be – black hole x-ray binary problem. trying to understand the reasons for which we do not observe be – black hole xrbs, belczyński & zió lkowski (2009) carried out stellar population synthesis calculations aimed at estimating the ratio of neutron star to black hole be xrbs, expected on the basis of the stellar evolution theory. the results of their calculations predict that for our galaxy the expected ratio of be x-ray binaries with neutron stars to the ones with black holes fns/bh should be, most likely, equal ∼ 54. since we know 48 neutron star be systems in the galaxy, then it comes out that the expected number of black hole systems should be just one. the observed number (zero) is consistent with this prediction. zió lkowski & belczyński (2010) carried out also preliminary stellar population synthesis calculations for the magellanic clouds. this time, the result was that the expected ratio fns/bh for magellanic clouds be xrbs should be equal ∼ 9. since we know 71 neutron star systems in the magellanic , then the expected number of black hole systems should be about 8. this time, the observed number (zero) is not consistent with the prediction. the most likely reason for this discrepancy is a different history of the star formation rate in the magellanic clouds, with respect to the galaxy. the calculations mentioned above used the galactic scenario for the star formation history. the new, more realistic calculations for magellanic clouds are necessary. 3 new be/x-ray binaries zió lkowski & belczyński (2011) published a full list of the 170 be/x-ray binaries known in 2011 (72 in the galaxy and 98 in the magellanic clouds). table 1 of this paper contains the list of the 14 new objects belonging to this class which were discovered in meantime (2 of them lie in the galaxy and 12 − in the magellanic clouds). this list contains two interesting systems (added after the end of our conference). one of them is system with the longest orbital period (swift j010745.0-722740 with porb ≈ 1180 days). the second is the first long sought be/black hole x-ray binary (agl j2241+4454). 4 candidates proposed for be/black hole x-ray binaries 4.1 unsuccessful candidates 4.1.1 ls i +610 303 this binary of the orbital period 26.5 d consists of a b0ve star and a compact object the nature of which is not firmly established yet. no pulsations were detected in the x-ray emission. the radio emission also does not show pulses. this might indicate a black hole rather than a neutron star. however, in many respects, the system reminds another x-ray and radio binary − psr b1259-63, the nature of which is already well understood. psr b1259-63 belongs to the category of so called colliding winds binaries. this wide (3.4 y orbital period) binary consists of an o9.5ve star and a neutron star which is observed as a fast (pspin = 47.7 ms) radio pulsar. neutron star moves along a very eccentric (e = 0.97) orbit. the system is also a source of tev emission (in fact, it was the first binary detected in this energy range). in all three energy ranges (radio, x-ray and tev) the emission is modulated with the orbital period. the binary was most recently extensively discussed by dubus (2013). after reviewing both the observations and the numerical simulation modeling, he strongly supports the much earlier idea of tavani et al. (1994) that both x-ray and tev emission are due to collision of two winds: stellar wind from be star and pulsar wind 176 be/x-ray binaries with black holes in the galaxy and in the magellanic clouds table 1: new be x-ray binariesa porb pspin l b x,max spectral name [d] [s] [erg/s] type refc xmmu j004814.0-732204 11.866 4.2 × 1036 b1.5-2.5ve 1 xmmu j005011.2-730026 29.9 214 5.0 × 1034 be 2,3,4 igr j00569-7226 17 5.05 5.5 × 1037 b0.5e 5,6,7,8 cxou j005758.4-721620 40.03 7.918 4 × 1035 be 9,10 2xmm j010247.4-720449 490 ? 521.4 2.8 × 1035 be 11,12 swift j010745.0-722740 1180 be 13 igr j01217-7257 84 ∼ 1 × 1037 be ? 14 cxo j012745.97-733256.5 656 1062 6.3 × 1035 b0-0.5iiie 15,16,17 igr j0154-7253 36.3 11.483 2.5 × 1037 o9.5-b0iv-ve 18 igr j015712-7259 35.1 11.6 4 × 1036 be 19,20,21 swift j04558.9-702001 5.1 × 1035 be 22 swift j053041.9-665426 ∼ 415 9.7 × 1036 be 23,24,25 maxi j1932+091 be 26,27 agl j2241+4454 60.37 3.7 × 1031 b1.5-2iiie 28,29,30,31,32 asince this table is a natural supplement to tables 1 and 2 of zió lkowski & belczyński (2011), i should add following updates to these tables: swift j0513.4-6547 porb ≈ 27 d (coe et al. 2013d) rx j0520.5-6932 pspin = 8.035 s (vasilopoulos et al. 2013b), porb = 23.93 d (kuehnel et al. 2014) smc sxp707 pspin is unknown (707 s is an instrumental effect, maggi et al. 2013) bmaximum x-ray luminosity c (1) sturm et al. 2011a; (2) coe et al. 2011; (3) schmidtke & cowley 2011; (4) schmidtke et al. 2013a; (5) coe et al. 2013a; (6) kennea 2013; (7) schmidtke & cowley 2013a; (8) coe et al. 2013b; (9) israel et al. 2013; (10) schmidtke & cowley 2013b; (11) sturm et al. 2011b; (12) sturm et al. 2013; (13) maggi et al. 2014; (14) coe et al. 2014; (15) henault-brunet et al. 2012; (16) schmidtke et al. 2012a; (17) schmidtke et al. 2012b; (18) townsend et al. 2011; (19) coe et al. 2008; (20) bodaghee et al. 2009; (21) schmidtke et al. 2013b; (22) vasilopoulos et al. 2013a; (23) sturm et al. 2011c; (24) charles et al. 2011; (25) sturm et al. 2011d; (26) negoro et al. 2014; (27) itoh et al. 2014; (28) williams et al. 2010; (29) casares et al. 2012; (30) casares et al. 2014; (31) paredes & ribo 2014; (32) munar-adrover et al. 2014. from rapidly rotating neutron star. the shock formed as a result of the collision gives the rise to a non-thermal emission in a similar way as in the pulsar wind powered nebulae. the characteristic element of such model is elongated radio emission with position angle dependent on orbital phase. elongated emission is pointing from the vicinity of the pulsar in the direction opposite to the direction towards the be star. such ”comet tail” behavior of radio emission is clearly observed for psr b1259-63. at present, there are no significant doubts that the system is powered by the pulsar spindown and that the colliding winds model is, in general, a correct explanation of the properties of this binary. ls i +610 303 is, in many respects, similar to psr b1259-63. in particular, the spectral and timing properties of the emission (through all ranges of electro177 janusz zió lkowski magnetic spectrum) are very similar. also the ”comet tail” behavior of radio emission is clearly seen for ls i +610 303 (dhawan et al. 2006). as i mentioned earlier, no pulsations were detected and, therefore, the case is not so clear cut as for psr b1259-63. this system (ls i +610 303) is also discussed by dubus (2013) and he concludes that indirect evidence strongly supports the colliding winds case. the compact object is, most likely, a rapidly rotating pulsar. the lack of the observed pulsations might be explained as the result of the absorption of the radio emission by the circumstellar material. even in psr b1259-63 the pulsed radio emission disappears for about 40 days around periastron. ls i +610 303 is much more compact binary (orbital period is ∼ 50 times shorter) and the absorption is strong throughout all orbital cycle. since the compact component is, most likely, a neutron star, the system is not a good candidate for a be/black hole binary. 4.1.2 ls 5039 this binary of the orbital period 3.9 d consists of an o6.5((f)) star and a compact object the nature of which is still a subject of controversy. no pulsations were detected neither in the x-ray nor in the radio emission. both x-ray and tev emission are modulated with the orbital period. the radio emission is variable but the flux is not modulated with the orbital phase. the morphology of the radio emission (which has an elongated shape) is variable with the orbital phase but the positional angle, at first look, seems to be roughly constant (moldon et al. 2012). this would not be a typical ”comet tail” behavior observed in psr b1259-63 and in ls i +610 303. however, analyzing more subtle changes in radio morphology and modeling radio emission with colliding winds model, moldon et al. (2012) demonstrate that the observed radio emission is compatible with the presence of a young non-accreting pulsar in the system. the authors obtain the best fit for a relatively high inclination of the orbit (i ≈ 700). as they note, this value has deep implication for the estimate of the mass of the compact component. radial velocities of ls 5039 (the optical component) were measured by casares et al. (2005). assuming pseudo-synchronization at periastron, they got i ≈ 200 for the inclination of the orbit and mx = 4.0 ÷ 7.3 m� for the mass of the compact component. this value clearly indicates a black hole. however, using the same radial velocities with the inclination derived by moldon et al. (2012), one gets only mx = 1.3 ÷ 2.7 m� for the mass of the compact component. this value corresponds to a neutron star and is compatible with the observed variability of radio morphology. it seems, that at present we cannot firmly establish the nature of the compact component. the black hole still cannot be excluded but its presence became somewhat doubtful. even if the presence of a black hole in the system would be confirmed, it is clear that the stellar component is not a be star and therefore the system is not a good candidate for a be/black hole binary. 4.1.3 maxi j1836-194 this system was discovered as an x-ray source by maxi in august 2011 (negoro et al. 2011). further observations strongly suggested that the system contains an accreting black hole (spectral states, timing characteristics, ejection of a radio jet). however, to classify the system as a be/black hole binary, we need a second ingredient − namely a be component. initially, it seemed that it is indeed present in the system. cenko et al. (2011) after spectroscopical observations of the optical counterpart with the 8 m gemini south telescope concluded that it is a be star. this finding was questioned by rau et al. (2011), who noted that the archival search of kennea et al. (2011) indicates that the optical counterpart undergoes outbursts of at least 4-5 magnitudes which is very unlikely for a be star. subsequent spectroscopical observations were made by pakull & motch (2011) who used vlt ut1 telescope. they found that the optical spectrum is not that of a be star. in this way, maxi j1836-194 joined the ranks of unsuccessful candidates. 4.2 the first successful candidate 4.2.1 mwc 656 this system was first discovered by agile as a gammma-ray source agl j2241+4454 (williams et al. 2010). the discovery paper identified the optical counterpart as a be star hd 215227 (known also under the name mwc 656) and determined the orbital period as equal 60.37 d. the nature of the optical component (of spectral type b1.5-2iiie) permits no doubts − it is a be star. the nature of the compact component was not established until very recently. casares et al. (2012) measured the amplitude of the radial velocities of mwc 656 and its rotational broadening obtaining 41.7 ± 6.8 km/s and 346 ± 10 km/s respectively. next, they analyzed fe ii lines reflecting the the keplerian rotation in the excretion disc around be star and under some assumptions estimated the inclination of the orbit as i = 67 ÷ 800. finally, they estimated the mass of the compact component to be in the range 2.7 to 5.5 m�. this range suggested rather a black hole but a neutron star could not be excluded. taking into account the overall low precision of this estimate, the nature of the compact component remained undecided. authors made more observations and published their analysis in 178 be/x-ray binaries with black holes in the galaxy and in the magellanic clouds a most recent paper in nature (casares et al. 2014). this time they got two sets of radial velocities: they measured the accretion disc emission line he ii 4686 reflecting the orbital motion of the compact component and they improved the measurements of radial velocities of be star by using sharp fe ii lines of the equatorial excretion disc. as a result they could come to a firm conclusion that the compact component is a black hole of the mass 3.8 to 6.9 m�. in this way the first be/black hole binary was finally found! the authors noted that the system seems to be xray quiescent (lx < 10 32 erg/s). this led them to a general comment that due to lack of a solid surface and a very low mass transfer rate (leading to extremely long outburst recurrence periods), be binaries with black hole companions might be difficult to detect by conventional x-ray surveys. after completing their work, the authors requested and obtained observing time with xmm newton. they found that the system emits xrays, after all (munar-adrover et al. 2014), but at a very low level lx ≈ 3.7 × 1031 erg/s, similar to the quiescent luminosity of black hole x-ray novae. this result does not change their conclusions. 5 summary during our conference, the statistics of be x-ray binaries in our galaxy indicated that we know 72 such systems. among them, 48 contain a confirmed neutron star and none, so far, is known to contain a black hole. the stellar population synthesis calculations (belczyński & zió lkowski 2009) indicate that the expected ratio of be x-ray binaries with neutron stars to the ones with black holes, fns/bh is equal ∼ 54. this ratio implies that, if we know 48 neutron star systems, then the expected number of black hole systems should be just one. few months after the conference this one system was just found! it is a binary agl j2241+4454/mwc 656. the problem for magellanic clouds remains unsolved (8 expected black hole be x-ray binaries and none observed). acknowledgement this work was partially supported by the polish ministry of science and higher education grant n203 581240 and by the polish national science center project 2012/04/m/st9/00780. references [1] belczyński, k., zió lkowski, j.: 2009, apj 707, 870. doi:10.1088/0004-637x/707/2/870 [2] bodaghee, a., tomsick, j.a., rodriguez, j.: 2009, atel 2252. 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[55] zió lkowski, j., belczyński, k.: 2011, in frontier objects in astrophysics and particle physics (procs. of the vulcano workshop 2010), f. giovannelli & g. mannocchi (eds.), conference proceedings, italian physical society, editrice compositori, bologna, italy, 103, 511 (arxiv:1111.2330). 180 http://dx.doi.org/10.1007/s10509-010-0575-8 http://dx.doi.org/10.1086/187542 http://dx.doi.org/10.1088/2041-8205/723/1/l93 introduction properties of be/x-ray binaries new be/x-ray binaries candidates proposed for be/black hole x-ray binaries unsuccessful candidates ls i +610 303 ls 5039 maxi j1836-194 the first successful candidate mwc 656 summary acta polytechnica ctu proceedings doi:10.14311/app.2015.1.0034 acta polytechnica ctu proceedings 2:34–39, 2015 © czech technical university in prague, 2015 available online at http://ojs.cvut.cz/ojs/index.php/app on traversability cost evaluation from proprioceptive sensing for a crawling robot jakub mrva∗, martin stejskal, jan faigl dept. of computer science, czech technical university in prague, technická 2, 166 27 prague, czech republic ∗ corresponding author: jakub.mrva@fel.cvut.cz abstract. traversability characteristics of the robot working environment are crucial in planning an efficient path for a robot operating in rough unstructured areas. in the literature, approaches to wheeled or tracked robots can be found, but a relatively little attention is given to walking multi-legged robots. moreover, the existing approaches for terrain traversability assessment seem to be focused on gathering key features from a terrain model acquired from range data or camera image and only occasionally supplemented with proprioceptive sensing that expresses the interaction of the robot with the terrain. this paper addresses the problem of traversability cost evaluation based on proprioceptive sensing for a hexapod walking robot while optimizing different criteria. we present several methods of evaluating the robot-terrain interaction that can be used as a cost function for an assessment of the robot motion that can be utilized in high-level path-planning algorithms. keywords: terrain traversability, proprioceptive sensing, walking robot. 1. introduction exteroceptive sensing is widely analyzed and summarized in the survey [1], where authors conclude that the most favored method is to assess the traversability characteristics before actually driving the robot into the respective region. such an assessment is based on a terrain model created from exteroceptive sensors like laser range finder or stereo camera. however, the difficulties a robot has while traversing rough terrain (e.g., slippages, softness of the ground, energy consumption, etc.) cannot be foreseen in advance. one has to actually walk through the terrain to feel and measure the interaction between the robot and the ground using proprioceptive sensors like accelerometers, force and torque sensors, etc. according to the survey that provides a study of unmanned ground vehicles (ugv), the generic capability of a robot to negotiate the terrain is the most commonly called as the traversability (along with occasional terms like terrainability, trafficability, mobility, etc.). based on that, we follow this established term and refer to its value as to the traversability cost. although the cost is a common name, its meaning varies among researchers. regarding the mission the robot is requested to accomplish, we need to balance between the time or the distance traveled, the energy consumption, the stability, the danger along the path, etc. having a terrain model, a common method of optimizing the traversability cost is to prefer a smooth and obstacle-free path, but a usage of proprioceptive sensors as indicators of the expected cost does not have such a common direction. in this paper, we report on existing approaches for traversability assessment based not only on exteroceptive sensors but we rather focus on proprioceptive sensors. moreover, instead of wheeled robot, we consider legged robots. the rest of the paper is organized as follows. we identify the main components of the terrain traversability analysis in section 2. first, we provide a brief overview of the exteroceptive sensing based methods for a comparison. next, we follow with an overview of known approaches using proprioceptive sensing. we formulate the main objectives in the view of the terrain traversability evaluation in section 3 and propose traversability evaluation methods for walking robots while focused on utilizing proprioceptive sensors in section 4. finally, a discussion about possible future approaches is given in section 5. 2. related work a great progress in the field of autonomous, perception-based, off-road navigation in robotic unmanned ground vehicles (ugv) was influenced by the darpa learning applied to ground vehicles (lagr) program [2], which ran from 2004 until 2008. the challenge evoked several solutions [3–5] with different approaches of robot sensing which can be divided into two groups as: 1) exteroceptive; 2) and proprioceptive sensing. 2.1. exteroceptive sensing it is reasonable to have information about the terrain prior to traversing it. laser scanner and camera are examples of exteroceptive sensors mostly used in the case of acquiring terrain characteristics. the terrain scan can provide an elevation map for a proper foothold planning [6] and estimation of its traversability [7] using a laser scanner. characteristic features can be obtained from a far-field scan (color image, stereo camera, etc.) and classified based on a model learned from near-field [3, 4, 8] where more data are 34 http://dx.doi.org/10.14311/app.2015.1.0034 http://ojs.cvut.cz/ojs/index.php/app vol. 2/2015 available. with data source getting closer to the robot, we get more precise models of the terrain, but usage of only exteroceptive sensors does not provide much information about the real interaction between the robot and the terrain being traversed and thus the robot can hardly learn from its experience. 2.2. proprioceptive sensing on the other hand, the robot needs to receive a feedback from the terrain to utilize learning from experience. proprioceptive sensors measure the modalities of the terrain that affect the robot motion directly. in the path planning task, the robot plans according to estimated traversability costs. however, the estimation can change when the robot actually encounters the terrain, as it has been applied in the previous approaches [3, 4, 8] using wheeled robot with bumpers, wheel encoders for a slip measurement or imu with gyros and accelerometers for measuring the roughness of the terrain. for example, a tall-grass area can be estimated as a non-traversable using only a range measurement, but it turns out that the robot can go through it without a significant difficulty. the traversability cost assessment is closely related to the terrain classification (or discrimination) assuming that the terrains of the same kind provide the same conditions for traversing. based only on proprioceptive sensors, the classification can be done using accelerometers [9] (the faster the speed, the better the accuracy) or yaw-angle variations [10] on a wheeled robot, or using servo drives feedback on a hexapod walking robot [11, 12]. focusing now mainly on legged robots, an early analysis of terrain traversability for legged locomotion using active perception was proposed by krotkov back in 1990 [13] studying the terrain stiffness and surface friction upon foot contact on a planetary rover – the ambler robot. however, the final path of the ambler robot is computed on a grid terrain elevation map avoiding occluded cells [14], i.e., without the possible knowledge gained from the proprioceptive sensors. a more matured solution [15] utilizes a biologically inspired gait on a hexapod crawler. carrying a lot of proprioceptive sensors, the robot was able to negotiate small obstacles using few hard-coded reflexes which ensured quality footholds during the motion. however, the estimation of the traversability cost relies only on exteroceptive sensors and the shape of the terrain. regarding our focus, we consider the work of hoffmann et al. as the most related to our approach. they presented how different sensory modalities affect the accuracy of the terrain discrimination using a quadruped puppy robot [16]. they also studied the relationship between the gait used and the classifier accuracy [17] including using sequences of different gaits to get better results. 3. problem statement the terrain sensing and the evaluation of the traversability cost of the robot motion fits into the scope of path planning. assuming we have a map of the environment in terms of positions of untraversable obstacles or regions, we can build a weighted graph on top of the map. although we consider evaluation of the cost for path planning, the planning itself is out of the scope of this paper. therefore, we consider the planning problem can be solved using optimal planners like a* [18] or d* [19] on a graph with known weights, i.e., the costs on the edges. the path consists of a sequence of actions (the robot moves along an edge using a particular gait) that have some cost—the traversability cost—which can be computed from various sources based on different criteria and different tasks the robot is performing. in general, we can distinguish sources of information the robot receives (terrain-sensor modalities) and the outcomes of the resulting robot motion (cost modalities) as can be seen in fig. 1. proprioceptive exteroceptive sensors terrain gait control f(s) evaluation c cost terrain-sensor modalities cost modalities figure 1. schema of the robot perception module. the input modalities represent different sensor measurements of the terrain which are used in a function f to evaluate the traversability cost c. the cost is multimodal in general but usually only one modality is used for planning (e.g., the robot speed). we can imagine that a terrain can be described by several characteristic features: shape, consistency, temperature, friction, etc. which can be measured by various sensors. each such a feature can be viewed as a modality that is either directly measured or remains hidden. each sensor refers to different modalities and combined together, we get a set of observable modalities on the input side of the robot perception module generating a multidimensional vector of sensor data. in general, we can say that more modalities— and hence more information—about the terrain we have, the better the estimation of its traversability cost we can get. notice that some of the modalities may provide more useful information (regarding the cost evaluation) than others. the output side of the module is also multimodal although it is not as obvious because usually only one modality is used for the traversability cost estimation (e.g., the robot speed while traversing the terrain). 35 j. mrva, m. stejskal, j. faigl acta polytechnica ctu proceedings the main problem is the middle part (of the schema in fig. 1) that consists of an evaluation function f which transforms the sensory measurement into a single (albeit multimodal) cost. the methodology how to extract a single cost from a lot of data in time, frequency or other domain is not a straightforward task and hence results vary among researchers from the simple and obvious findings to more and more sophisticated methods. 3.1. terrain-sensor modalities the input modalities refer to the robot-terrain interaction from which we can gather the sensory data. however, we can gather such data only from the observable modalities. for example, without a camera image, we can hardly determine the color of the terrain. a (certainly not complete) list of terrain-sensor modalities should include following terrain features: • shape (surface) – the 3d shape of the terrain or obstacles is useful for precise motion planning and it is usually scanned by exteroceptive sensors but can also be inferred from an advanced touch sensor. • color – a color camera image of the terrain area can enhance the terrain discrimination (e.g., flat sand vs. concrete). • consistency – a gravel, snow or sand is not a solid terrain and its surface is usually changed during and after the robot motion while leaving trace or marks on such terrain. this can be measured by sensible touch sensors or by comparing the surface shape before and after the motion referring to the terrain changeability or granularity. • softness – differences between solid (e.g., concrete) and soft (e.g., carpet) terrains can affect greatly the robot performance. this modality is measurable only by proprioceptive (e.g., force) sensors. • compliance – a terrain can look stiff (e.g., grass or small branch) but does not resist to the robot motion and usually regains its shape afterwards; so, a compliant terrain is penetrable. this is also measurable only by proprioceptive sensors. • adhesion – an adhesive terrain offers better friction and hence better traction which is crucial to ensure stable footholds and permit high accelerations. a force sensor or slippage analysis is needed to measure the adhesion with respect to the robot leg shape and material. • temperature – in some cases, e.g., on a volcano, a robot might need to avoid hot surfaces in order not to damage itself. the temperature (scanned by a thermal camera) can also be used for a better estimate of terrain traversability by analysing the relations between the terrain granularity (compactness) and thermal transients [20]. as it is shown in fig. 1, the values of input modalities measured by exteroceptive sensors are (in general) dependent on the terrain being scanned and also on the gait that drives the robot. the sensor readings can be affected when using a different gait— e.g., a fast tripod gait causes shaky motion of the robot and the images taken can be blurred. similarly, when measuring by proprioceptive sensors, some of the terrain-sensor modalities can change its value depending on the robot motion. for example, the terrain can seem solid and adhesive while traversing slowly, and crumbling or slippery when accelerating (on gravel or ice). another example for a wheeled vehicle is a sand (desert) that can be traversable until a small change in the control is applied and then it become completely untraversable. focusing on proprioceptive sensors and legged robots, the most important terrain-sensor modalities are those which directly affect the robot motion and hence its performance: shape, consistency, softness, compliance, and adhesion. 3.2. cost modalities in general, the outcome (or the cost) of the robot motion does not have a single scalar value. for example, answering a question: “how was the robot going?” with “5” looks like some information has been lost. instead, for example, we would like to know that it went fast but spent a lot of energy and hit several obstacles along the way. the crucial part is then to balance the trade-off between all of the possible outcomes—the cost modalities— of our interest. the outcomes (cost modalities) can be the following (again, we do not claim it is a complete list): • average speed (time) – the overall time and speed has to be evaluated relatively to the robot capabilities. either a reliable odometry or external motion capture system is needed for a proper evaluation. • energy consumption – a robot might have a limited energy capacity to fulfill its task, and therefore, it needs to care about the energy consumption (e.g., to switch to a more energy-efficient gait [21]). • maximal torque – in very rough terrains, the balance in exploiting all motors the same can be impaired and very high torque values of some motors can cause overheating or damage of servos and the robot itself. • uncertainty of localization – continuous robot body motion—if not smooth enough—can negatively affect the reliability of range sensors or camera-based visual localization. • stability risk – the robot posture can be close to its stability margins during the motion, which increases the risk of falling and should be counted in further path planning. • damage risk – a precise terrain analysis can unveil risky areas where the robot can be damaged (e.g., after a small slip), and therefore, these areas should be either avoided or at least considered. 36 vol. 2/2015 notice that the aforementioned cost modalities are dependent not only on the terrain the robot is traversing but also on the gait that is used to control the robot (except the risk of damage from a prior terrain analysis to avoid such a terrain). the traversability cost is usually evaluated using only a single modality when a robot is performing a simple task. however, if the robot has to perform a long-term mission and meanwhile learn from its own experience to achieve better results, a single modality is not enough and the other modalities have to be considered. a robot can then optimize its decisions in order to keep a high performance in a long period. 4. use case the proposed discussion of multiple modalities on both the terrain-sensor interface and the evaluation of the traversability cost can be applied on almost every mobile robot. however, each robot has different set of sensors and thus it can measure different terrain features and thus measure different cost modalities. here, we present an example—a use case—for a hexapod walking robot operating in a rough terrain with limited proprioceptive sensory data. the platform, used methodology, and testing scenarios are described in the following parts. in section 4.4, achieved results are presented. 4.1. used hexapod platform we used an affordable platform phantomx hexapod mark ii with an adaptive gait [22] that enables this robot to traverse uneven terrains and negotiate small obstacles. the gait is a periodic-based (in terms of alternating the legs in a given order) but alters and reacts on the underlying terrain surface. the robot utilizes its servo drives feedback for the tactile sensing of the ground and servo drives are also the only sensory information the robot has (i.e., the robot is technically blind). 4.2. methodology of cost evaluation following the set of possible traversability cost modalities listed in section 3.2 and the sensors available, we consider only the speed, energy (power) consumption, and maximal torque. the speed is given from the running time of the experiments and known distance of the traversed path. the energy consumption is estimated very roughly using several approximations. firstly, the instantaneous energy consumed (the power) is proportional to the drawn current assuming a constant voltage (p ∝ i). secondly, the current is proportional to the torque of the servomotor (i ∝ τ). thirdly, the torque is (according to the servo manufacturer) proportional to the servo drive controller position error e, (τ ∝ e). therefore, the power consumption (in a small discrete time step) can be inferred from the sum of absolute values of servo position errors p ∝ 18∑ i=1 |ei|. (1) for simplicity, we leave the units and scale because we only need to know the relative changes in the energy consumption under different conditions. finally, the maximal torque is computed similarly as τmax ∝ max e. (2) 4.3. testing scenarios the traversability cost evaluation was tested on three different terrains shown in fig. 2. the office floor is perfectly flat while the wooden blocks include obstacles with height about 5 cm (for comparison, the leg femur-tibia and tibia-foot links are 7 cm, resp. 13 cm long). the third terrain contains free wooden obstacles (2 cm high) that are not fixed to the floor and thus can be shifted during the robot traversal. the trajectory during experiments was equally long on all of the terrains and the robot was driven by adaptive gait [22] under 3 different configurations (pentapod, tetrapod and tripod). office floor movable obstacles wooden blocks figure 2. terrains traversable by the adaptive gait 4.4. results the measured experimental results are shown in table 1. as can be seen from the speed comparison, the robot is not slowed by obstacles and hence the used gait (in each of the three configurations) is very adaptive. nevertheless, its speed is slow even on the flat office floor and there can surely be found a faster gait for flat terrains which, however, may not be able to traverse other terrains. terrain office movable blocks gait 5/4/3 5/4/3 5/4/3 speed [mm/s] 9/15/22 9/15/22 9/15/22 work [e/mm] 45/28/19 52/32/23 47/28/21 power [e/s] 39/42/43 45/48/51 42/43/48 max torque 47/45/44 53/48/65 62/60/66 table 1. different cost modalities experimentally evaluated on different terrains for the same traveled distance using adaptive gait with different number of legs in support phase. three values in each cell stand for pentapod / tetrapod / tripod gait. 37 j. mrva, m. stejskal, j. faigl acta polytechnica ctu proceedings if we look at the traversability cost regarding different modality—the energy consumption, we can see that traversing the flat terrain is the least energy consuming. however, the other terrains have no more than about 10% increase in power consumption. traversing movable obstacles causes the highest energy consumption, which can be explained considering friction and leg configurations. when a leg slide sideways on a floating obstacle, a more momentum is created on the joints and thus more torque is needed to counteract this behavior. naturally, it is easier to walk with legs under the body than with legs straddled. another perspective of the evaluation is by comparing the maximal torque in the servo drives measured during the robot motion. we can see that the highest torque was measured when traversing the blocks. this indicates that such a terrain causes legs to be occasionally more loaded than others (e.g., after a small slip on the edge of a block), which is projected also into the average energy consumption (which is based on torque values). regarding the maximal torque values measured, we can also see that the tetrapod gait suffers less from high torque values (caused mainly by slippages), such that the impact after a slippage is not as big as for the pentapod or tripod gait. 5. conclusion we have shown that the evaluation of traversability cost, which is an important part needed for pathplanning, depends on the modality of sensory data as well as on the modality of the cost itself. each cost modality represents a different perspective of evaluating the robot performance and we show how sensory data can be transformed into the traversability cost estimation. while different situations need different cost modalities to be considered, in general, we need to find a trade-off between them to assess the cost more appropriately. we present a use case of the proposed idea in realexperimental evaluation with a hexapod walking robot traversing terrains with various difficulty. using only a single gait for all terrains might look appropriate according to the measured speed, but considering another perspective can unveil the potential risk of servo damage and switching to another gait (slower, but not the slowest) would be a more suitable solution. getting more into the problem of traversability cost estimation in the future work, we would like to take into account another modality—a gait modality. combined all modalities together, we can better model the perception of the robot which is a key factor in assessing the traversability costs to different terrain areas. moreover, the perception can be connected to learning and mapping between the terrain features and corresponding costs can be found automatically. acknowledgements the presented work has been supported by the czech science foundation (gačr) under research project no. 1509600y. the work of jakub mrva has been also supported by the grant agency of the czech technical university in prague, grant no. sgs15/208/ohk3/3t/13 references [1] p. papadakis. terrain traversability analysis methods for unmanned ground vehicles: a survey. engineering applications of artificial intelligence 26(4):1373–1385, 2013. doi:10.1016/j.engappai.2013.01.006. 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[22] j. mrva, j. faigl. tactile sensing with servo drives feedback only for blind hexapod walking robot. in proceedings of the 10th international workshop on robot motion and control (romoco), pp. 240–245. 2015. doi:10.1109/romoco.2015.7219742. 39 http://dx.doi.org/10.1177/027836499601500204 http://dx.doi.org/10.1016/j.robot.2014.07.006 http://dx.doi.org/10.1109/tssc.1968.300136 http://dx.doi.org/10.1109/icra.2015.7139750 http://dx.doi.org/10.1109/icra.2015.7139915 http://dx.doi.org/10.1109/romoco.2015.7219742 acta polytechnica ctu proceedings 2:34–39, 2015 1 introduction 2 related work 2.1 exteroceptive sensing 2.2 proprioceptive sensing 3 problem statement 3.1 terrain-sensor modalities 3.2 cost modalities 4 use case 4.1 used hexapod platform 4.2 methodology of cost evaluation 4.3 testing scenarios 4.4 results 5 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0089 acta polytechnica ctu proceedings 4:89–96, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app evaluation metrics applied to accident tolerant fuel cladding concepts for vver reactors martin ševečeka,b,∗, mojmír valachc a department of nuclear reactors, faculty of nuclear sciences and physical engineering, czech technical university in prague, v holešovičkách 2, praha 8, czech republic b alvel a.s., opletalova 37, praha 1, czech republic c ujv řež, a.s., department of severe accidents and thermomechanics, nuclear power and safety division, hlavní 130, řež, 250 68 husinec, czech republic ∗ corresponding author: martin.sevecek@fjfi.cvut.cz abstract. enhancing the accident tolerance of lwrs became a topic of high interest in many countries after the accidents at fukushima-daiichi. fuel systems that can tolerate a severe accident for a longer time period are referred as accident tolerant fuels (atf). development of a new atf fuel system requires evaluation, characterization and prioritization since many concepts have been investigated during the first development phase. for that reason, evaluation metrics have to be defined, constraints and attributes of each atf concept have to be studied and finally rating of concepts presented. this paper summarizes evaluation metrics for atf cladding with a focus on vver reactor types. fundamental attributes and evaluation baseline was defined together with illustrative scenarios of severe accidents for modeling purposes and differences between pwr design and vver design. keywords: nuclear fuel, accident tolerant fuel, evaluation metrics, lwr, vver. 1. introduction after the events at the fukushima daiichi npp, enhancing the accident tolerance of lwrs became a topic of high interest in many countries. the goal of accident tolerant fuel (atf) development is to develop alternative fuels to further enhance the safety, competitiveness, and economics of nuclear power. the world’s leader in atf development are the u.s., where in 2012, the congress directed to doe to: • develop enhanced fuels and cladding for lwrs to improve safety in the event of accidents in the reactor or spent fuel pools. • emphasize and fund activities aimed at the development and near-term qualification of meltdownresistant atf that would enhance the safety of present and future generations of lwrs. • create a plan for development of meltdown-resistant fuels leading to reactor testing and utilization by 2020. other countries (e.g. france, south korea, china) have been conducting their own atf research with similar goals. recently, russia introduced its plan for atf development with a focus on vver reactors. activities related to the atf development are also coordinated by international organizations as iaea or oecd/nea. the initial effort is focused on applications in operating lwr reactors (pwr, bwr, and vver). the goal set by the u.s. or south korea is to insert a lead test assembly (lta) into a commercial reactor by 2022. other countries including russia or the czech republic have not defined their specific goals, yet. this paper summarizes an evaluation methodology, proposed set of metrics, related tests and illustrative accident scenarios for the evaluation of vver atf cladding concepts. the metrics are based on the detailed evaluation approach proposed in the [1], with an emphasis given on the fuel cladding, vver reactors, and their departures from standard pwr reactor design. the evaluation methodology will be applied to assess each atf cladding concept relative to the current vver fuel system and will be described in details in future technical report. 1.1. atf definition atf fuels are according to the oecd/nea [2] defined as fuels that can tolerate a severe accident in the reactor core for a considerably longer time period than the current uo2 – zr alloy fuel system, while maintaining or improving the fuel performance during normal operations and operational transients. 1.2. coping time each concept will be evaluated besides other characteristics based on its ability to increase the “coping time” under severe accident conditions. where coping time is according to the oecd/nea defined as [2]: the time to significant loss of geometry such that the fuel can no longer be cooled or cannot be removed from the reactor. for each concept, there will be an analysis of failure modes and effects (fmea) performed to determine the onset of particular initiating event that leads to severe accident. however, the common baseline has to be set and general illustrative scenarios defined. 89 http://dx.doi.org/10.14311/ap.2016.4.0089 http://ojs.cvut.cz/ojs/index.php/app martin ševeček, mojmír valach acta polytechnica ctu proceedings 1.3. development strategy the strategy for atf development includes three phases: (1.) feasibility assessment and down-selection (laboratory experiments; code updates; assessment of economical, operational, safety, fuel cycle, and environmental impacts using evaluation metrics). (2.) development and qualification (atf fabrication; irradiation; safety basis testing; qualification and licensing). (3.) commercialization (technology transfer to industry). any new atf concept must comply with current safety and performance constraints, fuel cycle impacts or additional lwr design constraints. moreover, it’s attributes should be better in comparison with the current fuel system to achieve comparable or better performance in normal and accident conditions. for that reason, the quantitative metrics have to be developed and described in details taking into account also the requirements of utilities, fuel vendors, and regulatory bodies. 1.3.1. challenges in developing atf one of the main concerns is a definition of testing and qualification requirements. individual attributes are not equal, which requires concepts prioritization and determination which attributes to test and what metrics are needed to evaluate attribute compliance. if new atf concept should be accepted by utilities and vendors, it must be capable of integration into the current nuclear fuel cycle system (figure 1). the challenge is to get the best performance at each step of nuclear fuel cycle, and to understand how it affects other parts of the system. figure 1. atf shall not negatively (economically or technically) impact fuel cycle technologies. for that reason, integrated evaluation approach has to be adopted. 2. attributes of atf the attributes for mitigation of fuel failure during severe accidents provide fundamental guidance for atf evaluation. it may not be necessary to improve fuel system in all attributes. some attributes may provide substantial enhancement in accident tolerance, while others may provide only marginal benefits. the desired attributes highlight the performance of the fuel under normal and accident conditions. key attributes for a fuel system demonstrating enhanced accident tolerance include reduced steam reaction kinetics, lower hydrogen generation rate, and reduction of the initial stored energy in the core. the desired behaviors should be accomplished while maintaining or improving cladding thermo-mechanical properties, fuel-cladding interactions, and fission-product retention. in some cases, the described consequences of accident conditions may be concept-specific (e.g. hydrogen generation in the current zr alloy-uo2 fuel system). other atf concepts could present additional effects not expressed here (e.g. generation of co in case of sic) that must be considered in evaluation of the proposed system. a brief description of the desired attributes is provided in this section and summarized in figure 2. figure 2. major issues that need to be addressed in establishing accident tolerant fuel attributes. candidate fuel systems must first not harm, which means that the fuels must perform as well as or better than the current fuel system. moreover, the fuel system must additionally provide a comparable or improved response to aoos (anticipational operational occurences), dbas (design basis accidents) and bdbas (beyond design basis accident). 2.1. hydrogen generation rate hydrogen generation in the reactor core can lead to energetic explosions similar as in fukushima-daiichi accident. under a high-temperature steam environment, it is not possible to avoid hydrogen generation with standard zr-alloys. rapid oxidation of cladding results in free hydrogen generation. this exothermic reaction increases the cladding temperature, which further accelerates free hydrogen generation. a related issue is the diffusion of free hydrogen into the unoxidized portion of the cladding, resulting in enhanced embrittlement and potential cladding failure. the desired alternative is a cladding material that resists oxidation or reduces the rate of oxidation resulting in a slower hydrogen generation rate. materials with lower heat of oxidation are important due to the limitation of temperatures during an accident. materials that are less susceptible to hydrogen diffusion 90 vol. 4/2016 evaluation metrics for vver atf fuels may address the rapid embrittlement issue typical for standard zr alloys. 2.2. fission product retention cladding provides an important barrier between fission products and primary circuit. the potential release of fission products to the environment has to be avoided, therefore, retention within the fuel is of the highest importance. while total retention may not be possible, higher partial retention would be a substantial improvement. the desired improvement is to prevent melting or dispersion of the fuel by utilization of hightemperature/strength cladding materials that would retain cladding integrity beyond the current limitations of zr-alloy cladding. 2.3. cladding reaction with steam when cladding is exposed to steam at high temperature, multiple issues need to be considered: hightemperature steam interaction with zr-alloy cladding, an exothermic oxidation reaction, hydrogen generation, and degeneration of the structural integrity of the cladding. atf materials should demonstrate enhanced tolerance to radiation and oxidation under hightemperature exposure while specifically considering mechanical strength and structural integrity at the end of life and when exposed to high-temperature steam for an extended duration. 2.4. fuel-cladding interactions pellet cladding chemical interactions (pcci), pellet cladding mechanical interactions (pcmi) and fuel heating are important properties that must be understood during normal operation and accident conditions for all new fuel concepts. the desired design option is to develop fuels with reduced pcci and pcmi, with lower operating temperatures relative to the zr alloy-uo2 system, with higher melting point, and with structural integrity at high temperatures. chemical and physical compatibility of cladding and fuel for all proposed concepts must be ensured. 3. constraints on development of enhanced atf except for a few rare events, the current uo2-zr fuel system meets all performance and safety requirements while keeping nuclear energy economically competitive. any new atf concept should be compliant with and evaluated against current design, operational, economic, and safety requirements. fuel cycle considerations must also be considered, especially for concepts that represent a significant departure from the current technology. a brief summary of the constraints is illustrated in figure 3. figure 3. constraints on new fuel designs. the main constraints include: • potential atf geometry deviations and compatibility with co-resident fuel. • evaluation of all potential accident condition parameters, including: . fmea completion to ensure that any potential atf operating vulnerabilities are recognized and mitigated as possible. . completion of a § 50.59-like process (changes, tests and experiments section from the 10cfr50 regulation) to illuminate any necessary licensing or logistical preparations for operation of the atf in a commercial lwr. the process and its guidance is described in [3]. • quantifying the recommended minimum additional coping time to be provided by the atf. • the need to include a metric to address those concepts that require enrichment greater than 5 wt%. • the mission of the utility is not to test new fuels. any atf operation must necessarily address and minimize impacts on the utility by ensuring full compatibility with co-resident fuel and reactor components, and by limiting perturbations of the normal operation of the plant. • nuclear industry has optimized the current zr-uo2 which represents a large financial investment. the additional cost of atf must be relatively low. at present, it is not feasible to determine quantitative results. 3.1. backward compatibility atf concepts must be suitable for use in existing lwrs. longer term concepts may be considered in conjunction with the near-term focus. proposed fuel concepts should not require plant modifications to the current reactors. they should be compatible with existing fuel handling equipment, fuel rod or assembly geometry, and co-resident fuel in existing lwrs. 3.2. impact on operations a concept must maintain or extend plant operating cycles, reactor power output, and reactor control. re91 martin ševeček, mojmír valach acta polytechnica ctu proceedings figure 4. a diagram that describes an approach to assessing enhanced accident attributes of a proposed fuel design. ducing the availability or power output would be unacceptable to utilities. to maintain current operation, some of the fuel system concepts would require higher fuel enrichment. while the impact of higher enrichment is fairly well understood from a technical perspective, regulatory and safeguard issues have to be addressed. 3.3. economic impacts any atf concept is likely to be more expensive than the current uo2-zr system, at least initially. fuels that require higher enrichment are especially likely to cost more. increased enrichment could additionally require modifications to fuel fabrication facilities. therefore, it is important to carefully assess the economic impact of the new technology and to determine how much additional fuel cost the utilities will accept. on the other hand, increasing burnup and power densities (power upgrades) could reduce or mitigate the negative economical impacts. 3.4. safety impact performance of a new fuel system will be compared to the performance of the uo2-zr alloy system which defines a baseline for the atf evaluation. the performance must be shown to be similar to or better than of the current system. 3.5. fuel cycle impact atf must adhere to regulations and policies, for both the fuel fabrication facility and the operating plant, with respect to technical, regulatory, equipment, and fuel performance considerations. a new atf system could also have an impact on the back-end of the fuel cycle. the storage, repository performance of the fuel or possibility of reprocessing must not be degraded. 4. metrics and related testing the metrics determine a clear technical methodology for evaluation that can be used to rank two or more concepts. the approach has to be simplified due to the complex multiphysics behavior of nuclear fuel and the large set of performance requirements that must be met. dozens types of atf cladding ranging from metallic to fully ceramic cladding each with very different material properties and behavior has been studied. each concept has to be evaluated and compared to determine the concept viability, and prioritize resources to obtain the cladding with the best compromise in terms of properties and behavior in both nominal and accidental conditions. 4.1. desired cladding properties, behavior and performance the cladding of a nuclear fuel has three fundamental functions: (1.) confinement of fissile material and fission products within the fuel rod. (2.) not affecting chain reaction (low absorption cross section). (3.) efficient heat transfer from the fuel to the coolant. to fulfill these three functions a new ideal cladding material needs to have certain properties and exhibit specific behaviors as follows: fabricability: • compatibility with fabrication facilities; material availability; welding/sealing behavior. thermo-mechanical behavior: • high thermal conductivity, strength, and ductility; leak-tightness throughout plant operation (impermeable to fission gas and fission products). corrosion behavior: • high corrosion resistance in vver reactor environment; high corrosion resistance in high temperature steam; low h2 production and low hydrogen embrittlement. chemical compatibility: • compatibility with fuel, coolant, and the fuel assembly components; no eutectic. irradiation behavior: • stable or predictable thermal, mechanical, and corrosion behavior under irradiation; dimensional stability; low irradiation embrittlement; low thermal neutron absorption cross-section; low activation. back-end behavior: • low tritium permeation; no impact on the reprocessing; satisfactory storage behavior. new atf cladding material will not exhibit all presented ideal properties. for that reason, a baseline for evaluation and assessment of various concepts has 92 vol. 4/2016 evaluation metrics for vver atf fuels to be established. this corresponds to the metrics described further. 4.2. metrics definition the metrics are divided into three sections: (1.) material properties inherent to the material. (2.) material behavior observed through standardized tests or experiments for normal and accident conditions. (3.) material behavior observed through standardized tests for accident conditions. note that, material behavior can depend also on particular reactor type and its performance. 4.2.1. inherent material properties the inherent material properties that should be compared are mainly thermal and neutronic properties: high melting/sublimation temperature; high thermal conductivity; optimal heat capacity/enthalpy; optimal thermal expansion coefficient to ensure dimensional stability; thermal neutron absorption crosssection. there are other concept-specific inherent properties. for example in case of multi-layer coating, compatibility between different materials has to be ensured. additionally, eutectic temperature might decrease the melting temperature of the cladding so it should be avoided. 4.2.2. normal operation behavior in contrast with the inherent physical properties described above, the behavior of the materials has to be assessed through tests with standardized conditions to enable direct comparison and evaluation of the various concepts. long-term corrosion tests are proposed to evaluate the corrosion behavior in normal operating conditions in a different chemical environment. an acceleration of kinetics is likely under irradiation which has to be also investigated. the important data to assess are: visual inspection – potential delamination of coatings; corrosion kinetics – weight change, oxide thickness; water chemistry analyzes; hydrogen pick-up. mechanical tests suggested for mechanical behavior evaluation are: tensile tests; internal pressure creep tests; internal pressure burst tests; leaktightness behavior; fatigue tests – thermal cycling, internal pressure load cycling. 4.2.3. accident conditions behavior first, tests simulating normal operating conditions will be performed. later, the material performance under selected accident conditions will be assessed. specific tests selected to determine accident performance of candidate cladding materials are series of standard tests simulating ria and loca conditions. for example: high temperature oxidation, isothermal internal pressure burst tests at multiple temperature set points, dynamic internal pressure burst tests with increasing temperature until cladding failure, rapid burst test with increasing temperature until cladding failure, tensile tests at elevated temperature, internal burst tests, tensile test at elevated temperature etc. 5. evaluation baseline at the beginning of the evaluation, it is recommended to establish a common understanding of the current state-of-the-art cladding. key characteristics and performance criteria were identified, with quantitative values assigned where possible. all of the parameters are zr alloy specific, some are reactor design and reactor performance (chemistry regime, power changes) specific. however, there can be common baseline for most of lwrs defined with only small uncertainties. 5.1. baseline for normal operating conditions generally accepted characteristics of zr-based alloys during normal operations are [4]: • fission product retention: 10−6 pin failure rate, radiation tolerant: up to 10 dpa, 1022 n/cm2 fast (>1 mev), dimensional stability: max 1 % axial elongation. • thermal conductivity λ: ∼17.4 w/m k, heat capacity cp: 293 j/kg k, heat transfer coefficients: coolant/cladding coupling (reactor design dependent). • corrosion performance: oxide growth ≤100 µm, hydrogen content ≤800 wppm, maximum limit on corrosion product buildup (linked to water chemistry specifications), galvanic corrosion limits, requirement for all corrosion products to be harmless to other system components. • tritium release into coolant: limit for tritium permeation not established, but “low” permeation is desirable, <10 % of the total tritium generated permeating into the coolant could be reasonable. vendors assume that less than 2 % of this tritium permeates through zr-based alloy cladding to the coolant (e.g. for ss-316, the permeation could be as high as 90 %). • hydrogen pick-up: typical for zr-based alloys, depends on corrosion level and varies with alloy from 100 ppm to 800 ppm (600 ppm should be acceptable under normal operating conditions). • mechanical criteria – ductility: currently 1 % minimum of total elongation; modern alloys show uniform elongation up to 5 % and total elongation up to 20 %; yield strength: current as-fabricated alloys, ∼150–400 mpa at 340 ◦c, ultimate tensile strength: 437 mpa at room temperature (zry-4), creep strain limit: 0.66 % (zry-4); resistance to crack propagation (fracture toughness). 93 martin ševeček, mojmír valach acta polytechnica ctu proceedings • fretting wear: <10 % of wall thickness; debris wear resistance; fabricable: weldable, thin-walled tube that maintains hermeticity. • pci interaction behavior: includes consideration of fission product interactions (possible scc), oxide layer on the inside of the clad (∼10 micron). • burnup limit: peak rod average: 62 mwd/kg u. 5.2. baseline for accident conditions the accident performance of zr-based cladding differs from the normal operation behavior. the main current metric for fuel system performance is that the cladding should maintain post-quench ductility by limiting the peak cladding temperature to 1204 ◦c and the maximum oxidation level to 17 % of the cladding wall thickness. it has been shown that the temperature and oxidation limits result in maintaining ductility for as-fabricated and very low burnup cladding. changes underway nrc regulatory guides and iaea guidelines will limit the oxidation level for dbas based on pre-dba hydrogen content (e.g. 4 % for 600 wppm hydrogen content) [5]. under beyond bdba loca conditions, cladding will be subjected to higher temperatures and oxidation levels. zr-based cladding alloys would experience longer times at increasing temperatures, higher oxidation rates, higher hydrogen release rates, and higher internal heat-of-oxidation rates. key performance measures for fuel behavior during accident conditions include coping time, behavior under elevated temperature conditions over long periods of time, and material oxidation [6]. the relevant temperatures for lwr severe accidents which result in the formation of liquid phases due to melting or chemical interactions are summarized in figure 5. • coping time under dba and bdba – lbloca conditions: from 50 to 300 sec at 1204 ◦c, sbloca conditions: ∼1 hr at ≤1050 ◦c. • elevated temperature issues – melting temperature: ∼1760 ◦c, temperature for eutectic formation (decrease of melting point, effects on other components), runaway oxidation – relates to heat of oxidation and acceleration of the oxidation process. • steam oxidation kinetics: oxidation rate and associated heat of oxidation, performance considered for time at elevated temperature and maximum temperature, breakaway oxidation may occur at elevated temperature (e.g. time at ∼1000 ◦c before breakaway oxidation occurs should be defined). • hydrogen production and release: <1 % zr-metal conversion. • maintain ductility following dba: most important accident performance criterion under licensing standards, regulation specifies at least 1 % ductility post-quench. • high temperature mechanical properties: balloon and burst are not currently defined as a fuel failure and are not limited in the licensing criteria; while only a fraction of rods burst, those that do typically burst at ∼800 ◦c; burst may occur at lower temperature with higher burnup due to higher internal pressure; burst is a ductile failure; size of ballooning and burst opening can be important to the results of the failure. • severing and/or shattering the cladding with subsequent fuel release are considered as a failure; creep performance must be also considered. • flow blockage under accident scenarios: required calculation for licensing, related to maintaining coolable geometry, flow blockage may result from swelling / ballooning of cladding. • thermal shock resistance: current cladding alloys do not shatter until they have reached high oxidation levels (>17 % of the cladding wall thickness consumed by oxidation), standard test applied to determine thermal shock performance. figure 5. lwr severe accident-relevant melting and chemical interaction temperatures which result in the formation of liquid phases [7]. the goals of evaluation are associated with defining requirements for down-selection among options during the feasibility assessment. it includes identification of important attributes for atf, understanding the common baseline for evaluation and map the merit of the attributes against potential operational or safety envelope benefits which are presented in the previous sections. it is also required to define safety analyzes including accident scenarios to quantify the target 94 vol. 4/2016 evaluation metrics for vver atf fuels values of particular atf fuel characteristics which are presented in the next section. 6. illustrative scenarios for atf evaluation two “bounding scenarios” for bdba evaluation are proposed. they cover wide range of severe accidents and at the same time are not too prescriptive and specific to a particular npp design. modeling of these scenarios will utilize one of the initiating events and will continue until a defined point of failure as defined in by the coping time definition (see section 1.2) for particular atf concept. the two scenarios applicable to vver reactors are: • long-term station blackout: high-pressure scenario; it will be calculated to the point of reactor pressure vessel failure. • large-break loca: low-pressure scenario with high decay power and limited availability of emergency cooling systems. the proposed scenarios are intended to provide bounding cases for fuel performance. each atf concept will be evaluated by fuel performance and system analysis codes in regard to the two illustrative scenarios. all atf evaluations should be allowed to progress to the point defined by the coping time definition with considering failure modes and effects for particular atf concept. the goal of the evaluation is to estimate the potential increase in coping time that is offered by the atf concept and to assess potential outcomes (e.g. fuel failure, coolability, cause of failure). following completion of bounding (most severe) analyzes, more detailed studies for these illustrative scenarios should be performed to develop a better understanding of the impact of additional variables as burnup, time after scram when core cooling is lost etc. 6.1. high pressure scenario – sbo the high-pressure sbo (station blackout) scenario for atf evaluation purposes for vver-1200 reactor is defined as: • loss of all sources of ac supply. • feed water supply is unavailable. • turbine isolation valve is activated, pressure in the steam generators (sgs) increases. • after exceeding the pressure limit the quick-acting pressure reducing system opens. • a failure of the quick-acting pressure reducing system is postulated – system stays open after pressure decreases. • loss of coolant in sgs leads to decay heat removal passive system failure. 6.2. low pressure scenario – lbloca + sbo the low-pressure scenario for atf evaluation purposes for vver reactors is defined as: • guillotine rupture of the primary circuit’s cold leg near the reactor inlet with sbo. • decay heat removal passive system (dhrs) and emergency core cooling system are in operation. operation of the hydro accumulators is assumed. the decay heat removal passive system and emergency core cooling system operate according to their design characteristics: eccs – up to water volume exhaustion, dhrs – up to the end of nitrogen absorption due to eccs operation. total operation time of these systems is approximately 24 h from the accident start. 7. vver fuel specifics the russian design pressurized water reactors called vver have specific differences in comparison with western pwrs. it is recommended to adopt all evaluation metrics with both illustrative scenarios presented above and define additional vver specific metrics. the main differences of the vver fuel and primary circuit include: • different water chemistry, • hexagonal geometry and arrangement of the fuel assemblies, • various cladding and fuel materials . e110 alloy, e635 alloy, . annular pellet, different dopants, only gd-based burnable absorbers, • assembly shroud – in case of vver-440 fuel assemblies, • different control rod materials. the main difference between vvers and pwrs can be observed in nominal operation due to different chemistry of primary circuit and different material composition. while pwrs use lioh for chemistry control, vvers use koh. this difference in chemistry regime leads to slightly different test conditions for normal operation corrosion tests. however, in accident conditions the performance of vvers is similar to that of pwrs. therefore it is not necessary to define additional vver-specific metrics for accident conditions. 8. conclusion to develop new accident tolerant fuel systems an evaluation and prioritization of atf concepts has to be performed at the end of the first development phase. the prioritization will allow researchers and decision makers to focus resources on most promising concepts. due to the complicated complex multiphysics behavior of nuclear fuel it is not possible to test all the required characteristics for each concept. for that 95 martin ševeček, mojmír valach acta polytechnica ctu proceedings reason, evaluation metrics have to be defined. a good starting point for atf evaluation establishment is a common understanding of the current state-of-the-art zr-cladding and its performance in bdba conditions. it is a baseline for evaluation and assessment of various concepts. evaluation of atf cladding must consider the complete fuel system (including fuel cycle) and should encompass all performance regimes for the fuel, including: fabrication, normal operation and aoos, dbas, bdbas, and used fuel storage, transportation, and disposition. there are numerous attributes within each regime that must be considered in evaluating the fuel system performance. key attributes for the cladding were discussed in sections 4 and 5, along with a summary of standard tests recommended for measuring specific properties and characterizing performance under the specified conditions. it may not be possible to improve the current state-of-the-art fuel system in all attributes and regimes, significant improvement in some of the key attributes may outweigh modest performance gains or modest vulnerabilities in other attributes. a detailed list of proposed attributes for evaluation is provided together with metrics and standardized tests. the attributes summarized in this document, can be used as a qualitative guide to assess the performance of candidate materials relative to the current state-of-the-art materials and relative to one another. evaluation baseline for common understanding of the current cladding was presented. to further determine the common understanding of atf performance two severe accidents were defined and described. based on the illustrative scenarios the coping time can be calculated and atf concepts evaluated. most of the metrics and attributes are generally applicable for all lwr reactors but some attributes as chemistry regime are reactor-specific and for that reason additional metrics and standardized test were defined for vver types of reactors. detailed more prescriptive approach for evaluation of atf cladding for vvers will be described in future technical report. acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs ohk4-008/16. references [1] s. m. bragg-sitton, j. carmack, f. goldner. evaluation metrics applied to accident tolerant fuels. tech. rep., idaho national laboratory (inl), 2014. [2] egatfl. light water reactor accident tolerant fuel: evaluation metrics and technology readiness level definition. tech. rep., oecd/nea, 2016. [3] usnrc. 1.187, regulatory guidance for implementation of 10cfr50.59. tech. rep., 2000. [4] m. billone, y. yan, t. burtseva, et al. cladding embrittlement during postulated loss-of-coolant accidents. tech. rep., argonne national laboratory (anl), 2008. [5] g. youinou, r. s. sen. enhanced accident tolerant fuels for lwrs–a preliminary systems analysis. tech rep 2013. [6] l. braase. enhanced accident tolerant lwr fuels national metrics workshop report. tech. rep., idaho national laboratory (inl), 2013. [7] p. hofmann. current knowledge on core degradation phenomena, a review. journal of nuclear materials 270(1):194–211, 1999. 96 acta polytechnica ctu proceedings 4:89–96, 2016 1 introduction 1.1 atf definition 1.2 coping time 1.3 development strategy 1.3.1 challenges in developing atf 2 attributes of atf 2.1 hydrogen generation rate 2.2 fission product retention 2.3 cladding reaction with steam 2.4 fuel-cladding interactions 3 constraints on development of enhanced atf 3.1 backward compatibility 3.2 impact on operations 3.3 economic impacts 3.4 safety impact 3.5 fuel cycle impact 4 metrics and related testing 4.1 desired cladding properties, behavior and performance 4.2 metrics definition 4.2.1 inherent material properties 4.2.2 normal operation behavior 4.2.3 accident conditions behavior 5 evaluation baseline 5.1 baseline for normal operating conditions 5.2 baseline for accident conditions 6 illustrative scenarios for atf evaluation 6.1 high pressure scenario – sbo 6.2 low pressure scenario – lbloca + sbo 7 vver fuel specifics 8 conclusion acknowledgements references 329 acta polytechnica ctu proceedings 1(1): 329–331, 2014 329 doi: 10.14311/app.2014.01.0329 concluding address franco giovannelli1 1inaf istituto di astrofisica e planetologia spaziali, area di ricerca di tor vergata, via del fosso del cavaliere, 100 i00177 roma, italy corresponding author: franco.giovannelli@iaps.inaf.it abstract before i officially conclude this workshop — far be it from me to attempt to compete with some concluding remarks already delivered at the meeting with various levels of passion by gennady bisnovatyi-kogan, giulio auriemma and sergio colafrancesco — i would like to comment on some of the highlights emerging from our fruitful week of discussions about multifrequency behaviour of high energy cosmic sources, without any pretension of completeness. keywords: photonic astrophysics particle astrophysics neutrino astrophysics. undoubtedly the advent of spacecraft gave a strong impulse to astronomy. starting roughly in the mid 1970s, almost the whole electromagnetic spectrum has been continuously surveyed in a large number of space experiments. astroparticle physics, a new field of physics, was born roughly twenty years ago by joining the efforts of the high energy astrophysics community and the particle physics community. during this relatively short period of time, astroparticle physics has developed strongly through studies of cosmic sources that are emitters of photons, charged particles, and neutrinos. some of the sources could produce gravitational waves that will probably be detectable in the near future with the new generation of ground–based gravitational experiments, such as ligo, and space– based experiments, such as lisa. meanwhile the large hadron collider (lhc) has been producing excellent results on the higgs boson, and in general about the study of pp collisions at tev energies, never obtained before in ground–based laboratories. these results, together with the results coming from the hubble space telescope (hst), with its deep survey of faraway objects of the universe, and with the vhe emission detected from a number of cosmic sources, both galactic and extragalactic, are witnesses of the validity of the big bang theory, described by the standard model. among many experimental and theoretical results discussed during this workshop, i would like to mention a couple of them that, in my opinion, will underpin the future of astroparticle physics. the first is the vhe sky at energies eγ > 100 gev. about 150 tev sources have been detected (see fig. 1, http://tevcat.uchicago.edu/). only 20 years ago, this sky was practically empty. it is expected that over the next decade the ongoing operation of fermi will be accompanied by observations with the current ground– based h.e.s.s., magic, veritas experiments and the planned cta and hawc experiments. data obtained in the very wide energy range from 100 mev to 1 pev will provide very deep insights into a number of problems of high-energy astrophysics and fundamental physics (see the review paper about tev astronomy by rieger, de oña-wilhelmi & aharonian, 2013). figure 1: tev sources catalog (http://tevcat.uchicago.edu/) the second is connected with the reionization epoch of the universe at z ≈ 20 − 30, when the first stars appeared (e.g. lamb & reichart, 2000; ciardi & loeb, 2000; bromm & loeb, 2002). therefore, quasars and grbs could be detectable up to roughly that epoch. type–ia supernovae (sne ia) have been detected up to z ≈ 1.7; the future jwst will be able to detect sn ia in the range 1.7< z < 3 (aldering et al., 2007). quasars have been detected at z = 6.419, 6.43 (fan et al., 2003), and z = 7.085 (mortlock et al., 2011), and grbs up to z = 9.4 (cucchiara et al., 2011). these detections support the theory fixing the reionization epoch to z ≈ 20 − 30 (see fig. 2). 329 http://dx.doi.org/10.14311/app.2014.01.0329 franco giovannelli after this workshop, the importance of multifrequency astrophysics and multienergy astro-particle physics once more appears evident. however, there are many problems in performing simultaneous multifrequency, multisite, multiinstrument, multiplatform and multienergy measurements due to: i) objective technological difficulties; ii) sharing common scientific objectives; iii) problems of scheduling and budgets; iv) political management of science. figure 2: sketch of the evolution of the universe from the epoch of recombination up to the present time. the epoch of reionization is also marked (after dai zhigao, nanjing university). the positions of the highest z quasars and grb are superimposed. during this fruitful workshop, we hope to have demonstrated once more the “vulcano theorem” enunciated in my concluding address in 1984: it is possible to develop science seriously even if smiling. and finally, i would like to conclude with a few wonderful words of dr daisaku ikeda (2001)–president of soka gakkai international (sgi)–reported in the booklet for today and tomorrow the thought for may 30th: “whoever has many friends has greater opportunities for growth. in this way, one both makes society a better place, and lives happier and more satisfied. in all cases, human relations, interpersonal interaction and communication are of vital importance. we must establish and nurture friendship and contacts with many people, both in our environment, and in society in general. in this manner our life will open up and will flourish”. we could go back to early childhood when we were like the little prince. one sees clearly only with the heart. what is essential is invisible to the eye (from the little prince, by antoine de saint exupéry). acknowledgement it is my pleasure to thank: the scientific organizing committee (j.h. beall, th. boller, r. hudec, l. sabau-graziati, a. santangelo); the informatics operator (francesco reale alias figaro, the factotum) for helping with informatics and in solving all related problems during the preparation and development of the workshop, and at the registration desk in palermo; the executive secretary, daniela giovannini, who smoothed any inconveniences that occurred at the registration desk with professionalism and moreover with her nice smile; the directors of inaf-iaps, dcu&ce-inta, mpe, e-o. hulburt ctr-nrl, st. john’s college, iaa/university sand 1, ascr/ai ondřejov & the czech technical university in prague; the actresses lisa colosimo and flavia giovannelli; the violinist alessandro perpich and pianist valentina usai; the staff of the splendid hotel la torre. on behalf of the soc and loc, i am pleased to thank all participants and especially the speakers for their active contributions in updating this workshop with their talks, alive with their discussions, and friendly with their attitudes. my special thanks to three colleagues and friends (gennady bisnovatyikogan, giulio auriemma and sergio colafrancesco), who kindly took on the challenging task of making the concluding remarks of the workshop. finally, on behalf of all participants, i would like to express my warm thanks to the chef, mr daniele inzerillo, who prepared a large number of delicacies for us. i hope to meet all of you once again during our next palermo workshop. references [1] aldering, g. et al.: 2007, astropart. phys. 27, 213. doi:10.1016/j.astropartphys.2006.11.001 [2] bromm, v., loeb, a.: 2000, apj 575, 111. doi:10.1086/341189 [3] ciardi, b., loeb, a.: 2000, apj 540, 687. doi:10.1086/309384 [4] cucchiara, a. et al.: 2011, apj 736, 7. doi:10.1088/0004-637x/736/1/7 [5] fan, x. et al.: 2003, aj 125, 1649. 330 http://dx.doi.org/10.1016/j.astropartphys.2006.11.001 http://dx.doi.org/10.1086/341189 http://dx.doi.org/10.1086/309384 http://dx.doi.org/10.1088/0004-637x/736/1/7 concluding address [6] ikeda, d.: 2001, for today and tomorrow the thought of 30th may, edizioni esperia. [7] lamb, d.q., reichart, d.e.: 2000, aip conf. proc. 526, 658. doi:10.1063/1.1361618 [8] mortlock, d.j. et al.: 2011, nature 474,616. doi:10.1038/nature10159 [9] rieger, f.m., de oña-wilhelmi, e., aharonian, f.a.: 2013, front. phys. 8(6), 714. doi:10.1007/s11467-013-0344-6 331 http://dx.doi.org/10.1063/1.1361618 http://dx.doi.org/10.1038/nature10159 http://dx.doi.org/10.1007/s11467-013-0344-6 acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0076 acta polytechnica ctu proceedings 7:76–78, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app evaluation of the fatigue response of polyester yarns after the application of abrupt tension loads emilio luiz vieira louzadaa, ∗, carlos eduardo marcos guilhermea, b, felipe tempel stumpfa, b a programa de pos graduacao em engenharia mecanica, av. italia, km. 8, rio grande, brazil b policab, av. italia, km. 8, rio grande, brazil ∗ corresponding author: emilioluizlouzada@gmail.com abstract. the discovery of oil fields in deeper waters through out the last decades has led the oil industry to the necessity of replacing the mooring systems of offshore platforms from steel to synthetic cables. consequently, both the industry and the academy started to join forces in order to better understand the mechanical behavior of such materials when subjected to different service conditions. this work aims to assess the change in the fatigue life of polyester (pet) yarns if the material is submitted to an abrupt tension load prior to the application of the fatigue cycling. it was found that the fatigue life of the yarns tested are substantially reduced if the specimen is subjected to this kind of abrupt load in comparison to virgin samples. keywords: mooring lines, synthetic cables, offshore platforms. 1. introduction in the course of the last decades, offshore oil fields had been discovered along the brazilian shore in increasingly deeper waters throughout the years, which forced the oil industry to look for mooring systems to substitute the traditional systems based on steel cables. with the increase of the water depth in which the platform will operate, these systems tend to become potentially heavy, to the point where the mooring system weight itself could overcome the platform floating forces, leading to the sinking of the structure [1, 2]. in order to overcome this situation, the oil industry started to look for lighter mooring systems, which meant lighter materials, and the obvious choice was to invest in the development of mooring cables based on polymeric materials. however, there was a lack of understanding regarding these materials’ mechanical response when submitted to the service conditions of an offshore platform, which basically forced the engineers and academics to develop new research in the field [1–3]. in the present days, polyester (pet) is found to be a widely used material in applications such as anchoring cables. its good performance when submitted to cyclic loadings and the fact that its fatigue life is not affected by its submersion in salt water was the main reason it was the first material to be considered by the industry to such application [4]. however, if the maximum tension force is considerably high, the fatigue life of pet is dramastically reduced [5]. it was found that when operating between cyclic tension loads of 10% and 70% of the material’s tension resistance, samples failed after approximately 100,000 cycles [6]. if the maximum load is reduced to 60% of the material’s yarn break load (ybl), fatigue life increases to 1,000,000 cycles [5]. cyclic tension loads, however the main loads to which the cables are submitted during service, are not the only ones. during the installation of the anchoring cables, the launch of the torpedo anchor might eventually be submitted to a sudden stop, caused either by a failure of the launching system or even due to a stop because of heavy weather conditions, which would lead to abrupt axial tension loads in the synthetic cable being launched/installed. it is expected that this kind of sudden (high) tensile load affects the future mechanical behavior of the polymeric material, specially its fatigue life. this paper aims to correlate the change in the behavior of pet under cyclic loadings before and after it was submitted to abrupt axial loads. such study will be based on the material’s diagram relating number of cycles up to failure (log n) and the minimum tension load throughout the cyclic loading. 2. materials and methods the material used was polyester (pet). yarn samples were of 500 mm length with sandwich-type terminals (figure 1) [7]. prior to the application of the impact load, and throughout all the tests, the samples were left 24 hours at controlled temperature of 20±2°c and relative humidity of 65±4%. tension tests were carried out in an emic dl-2000 electromechanic universal testing machine in order to obtain the material’s ybl of virgin samples. using a 500 mm/ min displacement ramp, and ten specimen, it was found that this pet has an ybl of 165,44 n. 76 http://dx.doi.org/10.14311/app.2017.7.0076 http://ojs.cvut.cz/ojs/index.php/app vol. 7/2017 evaluation of the fatigue response of polyester yarns after the application of abrupt tension loads figure 1. sandwich-type terminals. figure 2. impact testing device. a set of virgin samples were taken to the impact testing device (figure 2) and submitted to an axial impact load corresponding to 10% of the material’s ybl. this load is achieved releasing a 1.7 kg dead weight from a 300 mm height, which transfers to the sample a potential energy of 5 j. the maximum elongation of the samples were measured using a marker attached to the bottom end of the yarns running along a ruler. after the impact test, each specimen is taken to an instron 8801 servohydraulic unversal testing machine to be submitted to force-controlled cyclic tension loads. a total of eight ranges of tension loads were used during the cyclic tests: 10-90%ybl, 2090%ybl, 30-90%ybl, 40-90%ybl, 50-90%ybl, 6090%ybl, 70-90%ybl, 80-90%ybl. the frequency used was 0.1 hz. the same fatigue procedure was applied to virgin samples, in order to compare their behavior with those damaged by the impact. each load range was applied to ten different samples (both virgin and damaged by the impact test), which means that a total of 160 yarn samples were used throughout the fatigue assessment. figure 3. cycles up to failure versus load ranges: superposed results. 3. results 3.1. effect of impact load on the fatigue life of pet figure 3 shows the number of cycles up to failure versus the minimum load of each load range in the case of virgin samples and damaged samples. regardless of the minimum load applied, the virgin samples have a longer fatigue life compared to the ones that were previously submitted to the abrupt tension load. table 1 sumarises the number of cycles up to failure for both cases for each load range. 3.2. effect of impact load on the yarn elongation together with the tests to determine the material’s ybl, data regarding the samples elongation in rupture were collected. that was made in order to compare the elongation in rupture of the samples submitted to the quasi-static tests performed on the universal testing machine with those obtained when the yarns fail due to impact only. it was found that a potential energy of 6.5 j, which corresponds to a tension load of 13% of the material’s ybl, is required to take the yarns to fail by impact only. a set of ten samples were submitted to this condition and a scale was used to get the yarn’s elongation in rupture. results comparing the elongation of the yarns during quasi-static test and dynamic test are shown in table 2. it was found that the samples tend to elongate considerably less in rupture when the load applied is abrupt (51.6 mm versus 57.5 mm of the quasi-static tests). 4. conclusions the present paper showed a preliminary assessment of the change in the fatigue life of polyester yarns when they are submitted to an abrupt axial load in comparison to the virgin material, and it was found that a sudden load of approximately 10% of the yarn break load is enought to lead the material to a considerable drop on its fatigue resistance. in terms of the yarn’s elongation in rupture, it was also found that quasi-static tests lead to a larger elongation, and it 77 e. l. v. louzada, c. e. m. guilherme, f. t. stumpf acta polytechnica ctu proceedings minimum load [% ybl] 10 20 30 40 50 60 70 80 virgin samples 50 73 108 127 127 141 153 75 after impact 30 34 41 48 57 71 68 44 table 1. number of cycles up to fatigue rupture. elongation in rupture quasi-static test 57.52 mm impact test 51.6 mm table 2. elongation in rupture for quasi-static and dynamic tests. can be explained basically by the viscoelastic behavior of the material, because during the impact load the material has no time to dissipate the potential energy in terms of heat energy, for example. it is also interesting to notice that a sudden load of only 13% of the material’s yarn break load is sufficient to lead to its rupture by impact, which is considerably low and could be used to warn the industry to take that into consideration when operating with mooring ropes made of polyester. references [1] c. j. m. del vecchio. light weight materials for deep water moorings. phd thesis 1992. [2] m. m. salama. lightweight materials for mooring lines of deepwater tension leg platforms. marine technology 21(3):234–241, 1984. [3] f. e. g. chimisso. the past, the present and the future of policab: the challenge of synthetic mooring ropes anchorages at pre-salt petroleum basin, in brazil. proceedings of the 10th youth symposium on experimental solid mechanics 2011. [4] r. r. rossi. cabos de poliester para ancoragem de plataformas de petroleo em aguas ultraprofundas, in portuguese. msc thesis 2002. [5] f. v. de carmago, c. e. m. guilherme, c. fragassa, a. pavlovic. cyclic stress analysis of polyester, aramid, polyethilene and liquid crystal polymer yarns. acta polytechnica 56(5):402–408, 2016. doi:10.14311/ap.2016.56.0402. [6] r. l. bosman. on the origin of heat build-up in polyester ropes. oceans96 mts/ieee prospects for the 21st century conference proceedings 1996. [7] d. pfarrius, e. duarte, f. e. g. chimisso. theoretical and experimental modeling of a socket sandwich for use in tension tests of synthetic ropes. proceedings of the vith youth symposium on experimental solid mechanics 2007. 78 http://dx.doi.org/10.14311/ap.2016.56.0402 acta polytechnica ctu proceedings 7:76–78, 2017 1 introduction 2 materials and methods 3 results 3.1 effect of impact load on the fatigue life of pet 3.2 effect of impact load on the yarn elongation 4 conclusions references acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0019 acta polytechnica ctu proceedings 4:19–21, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app radiation degradation of pcbs in sediments: comparison between two methods ľubica darážováa, ∗, andrea šagátováa, b, vladimír nečasa, marko fülöpb, bumsoo hanc a institute of nuclear and physical engineering, faculty of electrical engineering and information technology, slovak university of technology, ilkovičova 3, 812 19 bratislava, slovakia b university centre of electron accelerators, slovak medical university, ku kyselke 497, 911 06 trenčín, slovakia c eb tech co., ltd., daejeon 305-500, south korea ∗ corresponding author: lubica.darazova@stuba.sk abstract. polychlorinated biphenyls are toxic compounds which have accumulated in river sediments in eastern slovakia. bioaccumulation could cause even cancer. radiation degradation with electrons is new and perspective method to dechlorinate pcbs in sediment matrix. we tested the influence of two difference chemical pretreatments and electron irradiation on pcb contaminated sediments. keywords: pcbs, cuso4 · 5 h2o, electron beam, irradiation, sediment. 1. introduction polychlorinated biphenyls (pcbs) were used as additives for lubricants, adhesives in transformers and capacitors. chemical and physical properties are following: inertness, resistance to heat, non-flammability and high dielectric constant. pcbs were produced all over the world, but the production was banned in 1998. in slovakia the producer of pcbs was the chemical company chemko strážske ltd. probably, a part of produced pcbs was released to the environment through the strážske channel. over the years the sediments in surrounding area of the chemical factory (the laborec river and the zemplínska šírava lake) were contaminated with pcbs. the contamination in strážske channel exceeds 400-times the limit, which is allowed in slovakia for the sediment matrix. behavior of pcbs in the soil is strongly dependent on the basic soil properties. great attention has been devoted to the pcbs problem in soil contamination in united kingdom, usa, canada, japan and sweden. in slovakia, the monitoring of pcbs was carried out, but specific solution has not been given [1, 2]. irradiation of pcbs with electron beam has not been used very often and it is new technology, which could provide non-combustion method for pcbs degradation in soil sediments. water is a product produced during radiolysis and it is also presented in soil samples contaminated by pcbs. the presence of oxygen during the irradiation process has negative effect to radiolytic dechlorination of pcbs contaminated sediments. soil matrice is rich of organic and anorganic compounds. the dechlorination yields of pcbs were reduced, due to the organic content of the soil, which could compete with required free radiacal reactions [3]. due to the hydrophobicity of pcb, organic-rich sediments could be easily solubilized in water with the use of organic co-solvent. during the radiolysis, hydroxyl radicals react with pcbs through the addition to the phenyl rings to produce various isomeric pcb radicals (arcl(oh)) [4]. 2. pcbs degradation with electron beam 2.1. simple pcbs irradiation with electron beam electrons interact at high energies predominantly by elastic scattering by bound electrons. the chemical changes in pcbs occur through secondary reactions of radiolytic species with pcbs (fig. 1). hydrogen atoms and hydroxyl radicals react with pcbs via addition to the phenyl rings, producing pcbs radicals. pcbs that remain in association with or within the sediment phase may undergo dechlorination by thermalized electrons [5, 6]. figure 1. mechanism of radiation dechlorination of pcbs [5]. 19 http://dx.doi.org/10.14311/ap.2016.4.0019 http://ojs.cvut.cz/ojs/index.php/app ľ. darážová, a. šagátová, v. nečas et al. acta polytechnica ctu proceedings 2.2. chemical pretreatment and electron irradiation of pcbs due to the fact that pcbs are hydrophobic compounds, it is required to use co-solvent. very often is used as co-solvent propan-2-ol, which assist with the dechlorination process of pcbs. it is essential to stabilize the formed aromatic radicals, that´s why it is need the compound which could easily donate hydrogen atoms, which could stabilize the aromatic cycle. such compounds are alcohols. trifan et al. [7] tested the influence of propan-2-ol, ethanol, thf and electron beam irradiation for pcbs destruction in transformer oils and solutions (fig. 2). the best results in terms of degree of dechlorination and convertion of pcbs into biphenyls were obtained when propan-2-ol was used. also, naoh or koh are needed for basic environment, which prevent the oxidation of cl– ions to cl2 [7]. figure 2. influence of the different solvent type on dechlorination of pcbs (dose of irradiation = 40 kgy, [pcbs] = 32.6 g dm−3, naoh = 1.5 g, temp. 65 ◦c, cb = conversion of pcb into biphenyl, dd = degree of dechlorination) [7]. 3. methodology our study was aimed to compare our method of chemical pretreatment and electron irradiation of samples to results obtained by colleagues from south korea, who chemically pretreated and irradiated the samples of soil contaminated by pcb from slovakia. they used lower energy of electrons, 2.5 mev. after irradiation at eb-tech to a dose of 400 kgy the total concentration of pcbs in soil samples decreased by more than 53 % (fig. 3). 3.1. chemical pretreatment and electron irradiation of pcbs in our experiment we used 5 mev electrons and chemical pretreatment: cuso4 · 5 h2o and k2co3. 20 g of wet sediment was mixed with 2 g of k2co3, which provided stable ph ∼ 10 which is essential to keep cl– anions during reaction. our samples were rich figure 3. pcb concentration in soil sample as a function of dose irradiated by eb-tech vs. ucea trenčín. of water and water molecules are competitive compounds in the process of pcbs irradiation. due to this fact, 1 g of cuso4 · 5 h2o, as a water scavenger, was added to the sample. samples were irradiated in the test tubes by electron accelerator with scanning beam uelr-5-1s at the university centre of electron accelerators (ucea) of the smu in trenčín. we applied doses: 100, 300, 500 and 700 kgy to chemical pretreated samples. determination of pcb in sediment was done at the department of toxic organic pollutants of the smu in bratislava. isotopedillution method by using of 13c-labeled standard solution was used. about 0.15 g of dried homogenized irradiated sediment was soxhlet extracted with toluene (8 hours). an 1/50 aliquot of the extract was applied on multi-layer silica column (44 % sulphuric acid/potassium hydroxide/silver nitrate on activated silica gel). the pcb extract was carefully concentrated and after dilution coupled with high-resolution mass spectrometry (hrgc). the initial total content of pcbs was intended to 1842.69 ng g−1. 4. results and discussion fig. 3 shows that with increasing dose, the total amount of pcb congeners is decreasing at our experiments. at dose 700 kgy the pcb degradation falls to 46 %. from the fig. 4, it is seen, that pcbs groups are decreasing with increasing dose, but also the increase occurs, when higher chlorinated pcbs are converting to lower ones. for environmental purposes, it is required to degrade pcb with higher efficiency. our methodology (ucea) was less efficient than the methodology of eb-tech, as we reached only 46 % decrease of total amount of pcbs at 700 kgy in comparison to their 53 % decrease at 400 kgy. it could be caused by energy of electrons and also by the chemical pretreatment method. for this purpose we are going to find better chemical pretreatment methodology in our future research. 20 vol. 4/2016 radiation degradation of pcbs in sediments figure 4. concentration of pcb groups in soil samples chemically pretreated by cuso4 · 5 h2o and k2co3 as a function of increasing dose of electrons. acknowledgements the authors would like to thank for financial contribution to the research to the sut grant scheme for support of young researchers. references [1] i. danielovič, j. hecl, m. danilovič. soil contamination by pcbs on a regional scale: the case of strážske, slovakia. pol j environ stud 23(5):1547–1554, 2014. http://www.pjoes.com/pdf/23.5/pol.j. environ.stud.vol.23.no.5.1547-1554.pdf. [2] i. danielovič, j. hecl, r. mati. polychlórované bifenyly a ich obsah v životnom prostredí regiónu zemplín. informačná publikácia, košický samosprávny kraj, 2009. [3] a. singh, w. kremers. radiolytic dechlorination of polychlorinated biphenyls using alkaline 2-propanol solutions. radiation physics and chemistry 65:467–472, 2002. doi:10.1016/s0969-806x(02)00360-2. [4] d. c. schmelling, d. l. poster, m. chaychian, et al. degradation of polychlorinated biphenyls induced by ionizing radiation in aqueous micellar solutions. environmental science & technology 32(2):270–275, 1998. doi:10.1021/es9704601. [5] h. yuan, h. pan, j. shi, et al. kinetics and mechanisms of reactions for hydrated electron with chlorinated benzenes in aqueous solution. frontiers of environmental science & engineering 9(4):583–590, 2015. doi:10.1007/s11783-014-0691-8. [6] m. chaychian. radiation-induced dechlorination of pcbs and chlorinated pesticides and the destruction of the hazardous organic solvents in waste water. ph.d. thesis, university of maryland, 2007. [7] a. trifan, i. calinescu, d. martin. transformation of polychlorrinated biphenyls (pcbs) into non-hazardous products by electron beam treatment. rev roum chimie 60(10):1053–1055, 2009. 21 http://www.pjoes.com/pdf/23.5/pol.j.environ.stud.vol.23.no.5.1547-1554.pdf http://www.pjoes.com/pdf/23.5/pol.j.environ.stud.vol.23.no.5.1547-1554.pdf http://dx.doi.org/10.1016/s0969-806x(02)00360-2 http://dx.doi.org/10.1021/es9704601 http://dx.doi.org/10.1007/s11783-014-0691-8 acta polytechnica ctu proceedings 4:19–21, 2016 1 introduction 2 pcbs degradation with electron beam 2.1 simple pcbs irradiation with electron beam 2.2 chemical pretreatment and electron irradiation of pcbs 3 methodology 3.1 chemical pretreatment and electron irradiation of pcbs 4 results and discussion acknowledgements references 331 acta polytechnica ctu proceedings 2(1): 331–333, 2015 331 doi: 10.14311/app.2015.02.0331 the golden age of cataclysmic variables and related objects ii concluding address f. giovannelli1 1inaf istituto di astrofisica e planetologia spaziali, area di ricerca di tor vergata, via fosso del cavaliere, 100 i00177 roma, italy corresponding author: franco.giovannelli@iaps.inaf.it abstract before to conclude officially this workshop — far from me the idea to attempt some concluding remarks already dealt at the meeting with various burning by joseph patterson, mariko kato, dmitry bisikalo, and rené hudec —, i would like to comment few highlights coming out from our fruitful week of discussions about the golden age of cataclysmic variables and related objects ii, without any pretension of completeness. keywords: cataclysmic variables and related objects photonic astrophysics. undoubtedly the advent of spacecrafts gave a strong impulse to astronomy; starting roughly from middle 1970ies almost all the electromagnetic spectrum was continuously surveyed by the many space experiments. the cataclysmic variables (cvs) historically were the first systems for triggering the studies of the accretion disk around white dwarfs (wds) starting from the 1960’s with the schools of warsaw (poland) and cambridge (uk). however, they lost rapidly their primeval importance because of the advent of the first x-ray space experiments that, with their limited sensitivity, were mostly detecting x-ray binary systems (xrbs) that showed x-ray emissions abundantly over the thresholds of their detectors. this thanks to the presence of neutron stars or black holes as companions of the optical low–mass or high–mass stars. the x-ray emission of cvs is about 2–3 orders of magnitude lower than that of xrbs. thus the bulk of cvs observations was coming for long time from optical and uv regions, and sometimes from ir and seldom from radio bands. in the last decade results coming from the new generation satellites, especially in the hard x–ray and γ– ray regions, renewed the interest of scientific community about cvs. this, together with the new developments in searching for the progenitors of type ia supernovae was the main reason for an new explosion of interest about cvs. during this week of a deep discussion about cvs and related objects, mostly through the physical processes occurring inside those systems, has shown the most powerful way for a better and faster development of our knowledge of the cataclysmic processes, rather usual in the universe. among the many experimental and theoretical results discussed during this workshop, i would like to remark one important old idea developed by vladimir lipunov in 1980’s, that, in my opinion, will stress the future of cvs physics. figure 1: lipunov’s diagram for gravimagneti rotators calculated for 1 m� white dwarf. the positions of the polars am her and ae aqr, and the intermediate polars dq her, ei uma, and ss cyg are marked (after lipunov, 1987). this is the physical description of cvs as gravimagnetic rotators (lipunov, 1982, 1987). in this way the behaviour of cvs is completely determined by the spin period of the wd and the gravimagnetic parameter y = ṁ/µ2, where ṁ is the mass accretion rate onto 331 http://dx.doi.org/10.14311/app.2015.02.0331 f. giovannelli the gravimagnetic rotator having mass m, and µ is its magnetic moment. this is valid also for all the compact objects, like neutron stars and black holes. in this way, the behavior of cvs is completely independent of the optical phenomena which until now have been those which prevailed in cataloging such systems in different classes. figure 1 shows a part of lipunov’s diagram where polars (pcvs), intermediate polars (ipcvs) and non–magnetic cvs (nmcvs) are situated. as discussed by giovannelli & sabau-graziati (this workshop), it appears evident that the most suitable approach for studying cvs from a physical point of view is to consider them as gravimagnetic rotators. the detection of several sw sex systems having orbital periods inside the so-called ’period gap’ opens a new interesting problem about the continuity in the evolution of cvs. are the ipcvs and pcvs smoothly connected via the sw sex-like systems placed just in between? in order to fully understand the emission properties and evolution of cvs, the mass–transfer process needs to be clearly understood, especially magnetic mass transfer, as well as the properties of magnetic viscosity in the accretion discs around compact objects. consequently, the investigation on the magnetic field intensities in wds appears crucial in understanding the evolution of cvs systems, by which it is possible to generate classical novae (e.g., isern et al., 1997) and type-ia supernovae (e.g., isern et al., 1993). in those catastrophic processes the production of light and heavy elements, and then the knowledge of their abundances provides strong direct inputs for cosmological models and cosmic ray generation problems. after this workshop it appears evident, once more, the importance of multifrequency astrophysics. however, there are many problems in performing simultaneous multifrequency, multisite, multiinstrument, multiplatform and multienergy measurements due to: i) objective technological difficulties; ii) sharing common scientific objectives; iii) problems of scheduling and budgets; iv) politic management of science. during this fruitful workshop, we hope to have demonstrated once more the “vulcano theorem” enunciated in 1984 in my concluding address: it is possible to develop science seriously even if smiling. but, as you probably suspected, this workshop has been organized under ”peaceful surroundings”. therefore, i would like to mention and to support tolstoy’s philosophy. ”think again”, was written by tolstoy in 1904, at the beginning of the russian-japanese war. the human conditions described by tolstoy in the war are ”like spiders in a glass”. this is a pamphlet that is not only against carnage...: ”again the suffering that benefits nobody, again the lies, and again the universal process of stupidity, the turning of men into beasts...” ...but that is also against the racism and the hypocrisy of the educated: ”scholars ... deal extensively with the laws of the migration of peoples, the relationship between the white and the yellow race, between buddhism and christianity, and according to their deductions and considerations justify the killing of men ...”. ... it shows unequivocally the tolstoyan base of gandhi’s non-violence doctrine: ”i cannot act in any other way than god requires of me, and therefore, as a man, i cannot take part in any war, neither directly nor through a third party, nor by giving orders, nor by cooperating in any form, nor by encouraging doing it: i cannot, i will not and i will not do it.” and finally, i would like to conclude with few wonderful words of dr daisaku ikeda (2001) – president of the soka gakkai international (sgi) – reported in the booklet for today and tomorrow the thought of 30th may: “the one who has many friends has greater opportunities for growth. in this way, one both makes society a better place, and lives happier and more satisfied. in all cases, human relations, the inter-personal interaction and communication are of vital importance. we must establish and nurture friendship and contacts with many people, both in our environment, and in society in general. in this manner our life will open up and will flourish”. we could go back to early childhood when we were as the ”little prince”. one sees clearly only with the heart. what is essential is invisible to the eye. (from ”the little prince” by antoine de saint exupéry). acknowledgement it is my pleasure to thank: the scientific organizing committee (dmitry bisikalo, massimo della valle, rené hudec, lola sabaugraziati, giora shaviv, paula szkody); the informatic operator (francesco reale alias figaro, the factotum) for helping in informatics and solving all related problems during the preparation and development of the workshop and in palermo, at the registration desk; 332 the golden age of cataclysmic variables and related objects ii concluding address the executive secretary daniela giovannini, who smoothed all the inconveniences occurred at the registration desk with professionalism and moreover with her nice smile. daniela giovannini (biologist) said goodbye to everybody and left in hurry tuesday afternoon because of the start of an important experiment in her institute (cnr/ibcn) about “treatment with a new generation of adjuvant to enhance the human immune response against the major infections”; the directors of the inaf-istituto di astrofisica e planetologia spaziali, roma, italy; intadepartamento de cargas utiles y ciencias del espacio, madrid, spain; ras-institute of astronomy, moscow, russia; inaf-osservatorio astronomico di capodimente, napoli, italy; ascr/astronomical institute, ondřejov & czech technical university, prague, czech republic; technion-department of physics, haifa, israel; department of astronomy, university of washington, seattle, wa, usa; the actresses lisa colosimo and flavia giovannelli; the violinist alessandro perpich and pianist valentina usai; the staff of the splendid hotel la torre. on behalf of the soc and loc, i am pleased to thank all participants and especially the speakers for their active contributions in rendering this workshop updated with their talks, alive with their discussions, and friendly with their attitudes. a special thank to the three colleagues and friends (joseph patterson, dmitry bisikalo, mariko kato, and rené hudec) who kindly accepted the not easy task of making the concluding remarks of the workshop. finally, on behalf of all participants, i would like to express my warm tanks to the chef, mr daniele inzerillo, who prepared for us a large number of delicacies. i hope to meet all of you once again during our next palermo workshop. references [1] giovannelli, f., sabau-graziati, l.: 2015, this volume. [2] ikeda, d.: 2001, for today and tomorrow the thought of 30th may, edizioni esperia. [3] isern, j., hernanz, m., garćıa-berro, e.: 1993, in white dwarfs: advances in observation and theory, m.a. barstow (ed.), kluwer academic publ., dordrecht, holland, nato asi ser., c403, 139. [4] isern, j., hernanz, m., abia, c., josé, j.: 1997, in frontier objects in astrophysics and particle physics, f. giovannelli & g. mannocchi (eds.), italian physical society, editrice compositori, bologna, italy, 57, 113. [5] lipunov, v.m.: 1982, ap&ss, 85, 451-457. doi:10.1007/bf00653467 [6] lipunov, v.m.: 1987, ap&ss, 132, 1-51. doi:10.1007/bf00637779 333 http://dx.doi.org/10.1007/bf00653467 http://dx.doi.org/10.1007/bf00637779 acta polytechnica ctu proceedings doi:10.14311/app.2016.3.0047 acta polytechnica ctu proceedings 3:47–50, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app structural and mechanical characterization of deformed polymer using confocal raman microscopy and dsc birgit neitzel∗, florian aschermayer, milan kracalik, sabine hild johannes kepler university linz, institute of polymer science, altenberger straße 69 4040 linz, austria ∗ corresponding author: birgit.neitzel@jku.at abstract. polymers have various interesting properties, which depend largely on their inner structure. one way to influence the macroscopic behaviour is the deformation of the polymer chains, which effects the change in microstructure. for analyzing the microstructure of non-deformed and deformed polymer materials, raman spectroscopy as well as differential scanning calorimetry (dsc) were used. in the present study we compare the results for crystallinity measurements of deformed polymers using both methods in order to characterize the differences in micro-structure due to deformation. the study is ongoing, and we present the results of the first tests. keywords: raman spectroscopy, dsc, deformed polymers, crystallinity. 1. introduction deformed polymers are used in a wide range of applications because of their advanced mechanical properties. especially, materials with high strength gained increasing importance. in order to obtain high strength materials with tailored properties, the morphology can be controlled by varying process conditions. nevertheless, there is still a lack of knowledge in the correlation between mechanical properties, processing conditions and polymer morphology. the microscopic structure is claimed as a crucial factor, which determines the mechanical properties. furthermore, mechanical performance is influenced by the crystallinity and chain orientation of a polymer. therefore, characterization of these parameters is important for developing new tailored polymeric materials. wide angle x-ray diffractometry or differential scanning calorimetry (dsc) are common methods to determine the overall crystallinity of polymers. besides, also spectroscopic techniques such as raman spectroscopy [1–4] have shown to be suitable techniques to observe process-induced changes in crystallinity [5] and chain orientation [3]. raman spectroscopy uses polarized laser light to scan the surface of a sample. when changing the polarization angle of the filter, the intensities of certain structure-related peaks in the raman spectrum change as well, which allows conclusions on the polymer morphology. the degree of orientation can be determined from these peaks via a material specific orientation function fcz, such as equation (1) for pp [3]: fcz = 0.134 · ∑ i973∑ i998 − 0.182 (1) see [3] it reveals if the polymer chains are randomly distributed (fcz = 0), in deformation direction (fcz > 0) or perpendicular to it (fcz < 0). as fcz depends on the polarization filter’s orientation, the differences in multiple spectra must be considered when determining its value. crystallinity α is supposed to be independent of the filter’s orientation. its equation sums up peaks of multiple spectra to determine its value. αraman = ∑ i809∑ i809 + ∑ i841 (2) see [5] directed force is applied to the samples, this direction is designated mod (machine orientation direction). for each sample two raman spectra were recorded. the first with polarisation set to 0 degrees, emitting laser light perpendicular to the mod. the second spectrum is taken with the laser aligned parallel (i.e. 90 degrees) to the direction of the applied force. we took 40000 measurements of one sample per mod to see if mod-dependent peaks in the spectra can be identified. the relative intensities were determined by integration of raman bands. the crystallinity was then calculated via equation (2) from the sum-spectra of both measurement series, parallel and perpendicular to the deformation direction. the second method we used, was dsc. from dsc curves, first the enthalpy of the sample can be calculated and from these results the crystallinity as well. one goal of this study is to compare the crystallinity values, obtained from both methods, because each technique is best suited for different surface textures. if a sample has a very smooth surface, then confocal raman microscopy can be used; for more irregular surfaces, measuring with dsc is more suitable. to characterize the polymer sample also on a macro-structure level, mechanical tests are ongoing. 47 http://dx.doi.org/10.14311/app.2016.3.0047 http://ojs.cvut.cz/ojs/index.php/app b. neitzel, f. aschermayer, m. kracalik, s. hild acta polytechnica ctu proceedings the mechanical investigation focuses on tensile testing. with this method, tensile strength and fracture behaviour can be analysed. furthermore, hardness tests could give us important data to build up the comprehensive material overview of the polymer samples in microand macro structure scales. 2. experimental in this current study, raman measurements were done with a confocal raman microscope (alpha300r, witec gmbh, ulm, germany). for the measurements, polarized laser light of 532 nm and 16 mw power was used. a polarization direction of 0° or 90° was applied. the graphical analysis has been performed using an objective with 20x magnification. with this equipment and these settings, polypropylene (pp) samples with different degrees of deformation in one direction were analyzed. furthermore, polyethylene terephthalate (pet) samples were investigated. the second method in this study was dsc using the instrument perkin elmer, dsc 8000, usa . both, heating rate and cooling rate were set at 10 k/min. the pp samples were heated from 25°c to 250°c, and the pet samples from 25°c to 300°c. the crystallinity was calculated from the first heating curve, to detect the influence, caused by the deformation. 3. results and discussion 3.1. crystallinity determined by confocal raman microscopy in the case of pp, the peaks at 809 cm−1 and 841 cm−1 provide information about the crystallinity, whereas the peaks at 973 cm−1 and 998 cm−1 are important for determining the chain orientation. these peaks change their orientation depending on the polarization of the laser light. all results were calculated via integration of the peak areas (see figure 1). figure 1. material specific (pp) peaks for crystallinity and chain-orientation calculation. in figure 2, the crystallinity distribution of a nondeformed (figure 2a), deformed pp (figure 2b), and highly deformed pp (figure 2c) can be seen. the deformed sample has a higher degree of crystallinity compared to the non-deformed sample. thus, the random distribution of the polymer chains, in the nondeformed sample changed to a well-ordered structure in the highly deformed pp-sample. (a) . non-deformed pp (b) . deformed pp (c) . highly deformed pp figure 2. crystallinity α calculated from raman spectra. confocal raman microscopy can be a powerful tool for structural analysis of further polymeric materials, but in this case it is necessary to find material specific structure relevant peaks. two peaks could be interesting for the second investigated material, pet: 998 cm−1 and 1096 cm−1. [6] furthermore, material specific formulas for crystallinity and orientation are required. 3.2. crystallinity obtained by dsc a big advantage of differential dsc is that samples can be produced very easily and in the same way for every polymer. also the crystallinity can be calculated with the same equation. so far, pet has been analysed as a second polymeric material. in figure 3, the crystallinity, calculated from dsc measurements, can be seen. these first measurements show, that crystallinity increases with high deformation. in a next step, these results will be compared to the crystallinities, determined via confocal raman microscopy. macrostructure will be characterized via tensile test. the aim is to find a relationship between the microand 48 vol. 3/2016 structural and mechanical characterization of deformed polymer figure 3. crystallinity of pet samples, measured via dsc; (a) amorphous, (b) deformed and (c) highly deformed pet. macro-structure behaviour of different polymeric materials. 3.3. comparison of crystallinity calculated by confocal raman microscopy and dsc using both methods, dsc and raman spectroscopy, it was found that the crystallinity α (figure 4) increases with higher deformation, but they have different increase rates. figure 4. crystallinity α calculated from raman spectra and dsc measurements for a non-deformed, b deformed and c highly deformed polypropylene (pp). the highest crystallinity and best correlation between both measuring methods was observed in the most deformed polypropylene c. it was furthermore observed that deformation of pp causes changes in its microstructure. 3.4. orientation distribution obtained by confocal raman microscopy figure 5 shows that the orientation function for the polymer chains, determined via confocal raman microscopy, increases with higher deformation. the polymer chains realign under stress by positioning themselves parallel to each other, resulting in a restructured polymer. this effect was also observed for the crystallinity, which also increases. a denotes non-deformed pp without orientation, b deformed polymer with weak orientation and c highly deformed pp with a strong chain-orientation along the deformation direction, 0° perpendicular, 90° parallel to mod. figure 5. orientation function calculated from raman spectra and dsc measurements for a nondeformed, b deformed and c highly deformed polypropylene (pp). (a) . parallel to mod, nondeformed (b) . perpendicular to mod, nondeformed (c) . parallel to mod, deformed (d) . perpendicular to mod, deformed (e) . parallel to mod, highly deformed (f) . perpendicular to mod, highly deformed figure 6. orientation distribution of pp samples calculated from confocal raman spectra figure 6 shows the orientation distribution of the pp-samples at different mods and degrees of deformation. the scanned area was 200 µm× 200 µm. it can be seen that the deformation changes the chain orientation from a random distribution to a well-defined periodic arrangement for both laser polarizations. the random distribution of the chains in non-deformed pp changes to an uniform distribution in highly deformed pp. 49 b. neitzel, f. aschermayer, m. kracalik, s. hild acta polytechnica ctu proceedings 4. summary with the two selected methods, dsc and raman spectroscopy, crystallinities were calculated. it was shown that the crystallinity and orientation of the chains increase with the deformation of the polymer. the effect of increasing crystallinity caused by deformation was shown for both polymeric materials, pp and pet. 5. outlook in the current study, pp as one important polymeric material, was analyzed and the results show that confocal raman microscopy and differential scanning calorimetry are powerful tools for the characterization of the inner structure. because of the different chemical behavior of different polymers, it is necessary to find useful structure peaks and mathematical formulations for crystallinity and orientation. finding solutions for other polymer, like pet, will be a subject of future work. in addition, relationships between mechanical properties and the polymer morphology/structure on a microand macro-scale are currently under investigation. acknowledgements robert führicht for latex typesetting. references [1] j. martin, m. ponçot, p. bourson, et al. study of the crystalline phase orientation in uniaxially stretched polypropylene by raman spectroscopy: validation and use of a time-resolved measurement method. polymer engineering & science 51(8):1607–1616, 2011. doi:10.1002/pen.21944. [2] m. fernandez, j. merino, m. gobernado-mitre, j. pastor. molecular and lamellar orientation of α-and β-transcrystalline layers in polypropylene composites by polarized confocal micro-raman spectroscopy: raman imaging by static point illumination. applied spectroscopy 54(8):1105–1113, 2000. [3] j. rodríguez-cabello, j. merino, l. quintanilla, j. pastor. deformation-induced conformational changes in stretched samples of amorphous poly (ethylene terephthalate). journal of applied polymer science 62(11):1953–1964, 1996. [4] j. rodríguez-cabello, j. merino, t. jawhari, j. pastor. rheo-optical raman study of chain deformation in uniaxially stretched bulk isotactic polypropylene. journal of raman spectroscopy 27(6):463–467, 1996. [5] a. s. nielsen, d. batchelder, r. pyrz. estimation of crystallinity of isotactic polypropylene using raman spectroscopy. polymer 43(9):2671–2676, 2002. [6] s. yang, s. michielsen. orientation distribution functions obtained via polarized raman spectroscopy of poly (ethylene terephthalate) fibers. macromolecules 36(17):6484–6492, 2003. doi:10.1021/ma034486q. 50 http://dx.doi.org/10.1002/pen.21944 http://dx.doi.org/10.1021/ma034486q acta polytechnica ctu proceedings 3:47–50, 2016 1 introduction 2 experimental 3 results and discussion 3.1 crystallinity determined by confocal raman microscopy 3.2 crystallinity obtained by dsc 3.3 comparison of crystallinity calculated by confocal raman microscopy and dsc 3.4 orientation distribution obtained by confocal raman microscopy 4 summary 5 outlook acknowledgements references 235 acta polytechnica ctu proceedings 1(1): 235–239, 2014 235 doi: 10.14311/app.2014.01.0235 data analysis of globular cluster harris catalogue in view of the king models and their dynamical evolution. ii. observational evidences marco merafina1, daniele vitantoni1 1department of physics, university of rome la sapienza, piazzale aldo moro 2, i-00185 rome, italy corresponding author: marco.merafina@roma1.infn.it abstract we summarize some observational comparison concerning the features of globular clusters (gcs) population in connection to the evolution of king models. we also make a comparison with some extragalactic gcs systems, in order to underline the effects of the main body on the dynamical evolution. keywords: globular clusters gravothermal catastrophe king models thermodynamical stability. 1 introduction globular clusters (gcs), for their proprieties of symmetry and their high relaxation times, are important to test theories about thermodynamical stability of spherical self gravitating systems. the actual sample of is a mixture of various and not homogeneus gcs types. therefore, it is difficult to analyze properties of milky way (mw) gcs population in connection to core-collapse and gravotermal instability. the last version (2010) of harris gcs catalogue (see also harris, 1996) includes 157 objects. it was pointed out by van der bergh (2011) that harris catalogue could includes three not typical gcs, probably remnant cores of dsph galaxies: omega centauri, terzan 5 and ngc 6715 (m54). the harris catalogue also includes pcc gcs, namely gcs with collapsed cores that cannot be described by classical single mass king models profiles. zinn (1985) identifies two classes of gcs, respectively known as disk population (metal rich) and halo population (metal poor), distinct by the threshold value [fe/h] ' −0.8 (or, according to some authors, −0.75). recently bica et al. (2006) showed that the actual gcs population seems to have been contaminated by capture of smaller galaxies (and their possible gcs populations) during the milky way formation. possible evidences of extragalactic origin of some gcs are retrograde motion (compared to galactic disk motion) and unusual young absolute age. it seems that the original gcs population suffered deep and incisive processes of disruption (see aguilar et al., 1988; hut & djorgovski, 1992; gnedin & ostriker, 1997; mackey & gilmore, 2004), until almost 50% of original gcs are destroyed in the last hubble time. 2 discussion we start to consider the problem introduced by katz in the paper about thermodynamic stability in 1980. the study of the distribution of galactic gcs in terms of w0 (central gravitational potential) shows a peak value of 6.9. we should expect that the peak value coincides with the known stability critical value w0 = 7.4, due to the old age of mw gcs and the onset of the instability in the high w0 region. this problem had remained unsolved (fig.1). figure 1: distribution of galactic gcs at different k (katz, 1980). the quantity k is related to w0 (see merafina & vitantoni, part i). with the introduction of the effective potential (see merafina & vitantoni, part i) and including the additional term in the expression of the total energy, we can revise the katz study. the result is a very satisfactory coincidence of the observative peak value with the stability limit. we can also repeat the analysis on a more detailed and updated sample (using data of the harris gcs catalogue). the peak value, in the non-symmetric 235 http://dx.doi.org/10.14311/app.2014.01.0235 marco merafina, daniele vitantoni gaussian hypothesis, is exactly at w0 = 6.9 (fig.2). 0 2 4 6 8 10 12 14 0 2 4 6 8 10 12 14 figure 2: w0 distribution of pre-core-collapse mw gcs. for a better understanding of the evolution of a gcs population, we briefly analyze the role of environmental features. the effect of the distance from the galactic center (fig.3) is not clear at all. generally speaking, the more a gc lives near the galactic center, the more quickly it evolves towards the gravothermal catastrophe, being more affected by tidal forces of the galaxy. 2 4 6 8 10 12 14 1 10 100 r g c figure 3: galactocentric distance rgc in function of w0. the dashed lines represent the values at 3kpc and 30kpc. if we look at the gcs distribution in the [fe/h]-w0 plane (fig.4), we find no correlations between these two quantities. this means that the difference between halo and disk population, first introduced by zinn (1985) does not influence the dynamical features and the evolution. nevertheless, if we analyze the disk population (fig.5), this seems to be more dynamically evolved than the halo one. the w0 peak value is slightly larger for the disk population, mainly due to a lower rgc (in average) for this class of objects. on the other hand, it is well known that the tidal shocks played a more incisive role in the evolution of the disk gcs. 2 4 6 8 10 12 14 -2.5 -2.0 -1.5 -1.0 -0.5 0.0 figure 4: total metallicity [fe/h] in function of w0. 2 4 6 8 10 12 14 0 2 4 6 8 10 12 n w0 external halo internal halo figure 5: w0 distribution of astronomical populations. the contamination of the halo gcs with the extragalactic origin ones is suggested by the bimodality in the absolute age distribution, by the rotation in the plane of galactic disk, and by the age-metallicity dependence. the situation is shown in figs.6, 7 and 8. regarding the age-metallicity dependence, we can say that few objects are located out of the main well defined sequence of clusters (and probably have a different origin with respect to all the others). 236 data analysis of globular cluster harris catalogue... part ii figure 6: absolute age distribution for mw gcs. older gcs, presumably all native in the milky way, are evidenced in dark grey. figure 7: motion over the galactic plane for mw gcs (from dinescu et al., 1999). figure 8: age-metallicity behavior, using the absolute age values extimated by dotter et al. (2010). pcc clusters are indicated by grey squares, suspected pcc ones by white triangles. if we analyze the behavior of the central relaxation time trc in function of w0 (fig.9), we can note a linear decreasing, that indicates an increasing of evolutional speed towards the collapse. below the treshold value log trc = 8.0 there is a region of coexistence of pre-core collapse gcs with post-core collapse (pcc) ones. the behavior of core collapse time tcc (that is the remaining time before the collapse of the model), changes over the stability limit w0 = 6.9 (fig.10). 2 4 6 8 10 12 14 4 5 6 7 8 9 10 11 r gc < 3 kpc 3 kpc < r gc < 30 kpc r gc > 30 kpc lo g (t r c/ 1y r) w0 ngc6717 figure 9: behavior of trc in function of w0; three classes of cluster distances are represented. vertical dashed line represent the stability limit, whereas the horizontal one represent the trc critical value (cohn & hut, 1984) distinguishing pre-core-collapsed and postcore-collapsed objects. suspected pcc clusters are indicated by filled symbols. in the harris catalogue ngc6723 has been erroneously included among the suspected pcc clusters in place of ngc6717. 2 4 6 8 10 12 14 7 8 9 10 11 12 lo g (t c c/ 1y r) figure 10: core-collapse time in function to w0. the tcc values are estimated as in quinlan (1996). we also consider a comparison among four gcs populations, shown in fig.11: the lmc is the less evolved, as well as the smc and fornax systems (mackey & gilmore, 2003b, 2003c); it presents a peak value close to w0 = 4.3. for this kind of gcs population, the presence of a low massive main body allowed to preserve 237 marco merafina, daniele vitantoni more informations about primeval distribution features. on the contrary, ngc5128, whose main body is a giant elliptical galaxy, seems to be an evolved population with a maximum value up to the threshold value w0 = 6.9. finally, the m31 system is the most similar to our gcs population, with a main peak value around the stability limit and an extended tail in low-w0 region, that is produced by the presence of low evolutionary speed objects or extragalactic origin clusters. also the effects of disk and bulge shocking, realistically, concurred to the formation of the tail. 3 conclusions in order to analyze the dynamical evolution of king single mass gcs, we have analyzed the mw gcs population. the mw clusters w0 distribution presents a peak very close to the new stability limit w0 = 6.9 and a pronounced tail in the low-w0 region. 3 6 9 12 0 5 10 0 5 10 0 3 6 3 6 9 12 0 3 6 w0 mw n m31 lmc figure 11: comparison among w0 distributions of different gcs systems. in addition to mw system, m31, ngc5128 and lmc systems are reported. m31 data come (or are deduced) from barmby et al. (2007), ngc5128 ones from gòmez et al. (2005), lmc data from mackey & gilmore (2003a). mw histogram differences from fig.2 are given by a different binning choice. we can instead exclude a direct relation between astronomical gcs populations and dynamical evolution, except for a very weak increasing of w0 peak value for disk clusters. from the time-scales we can deduce that clusters with high w0 value have an higher collapsing speed. the gravothermal catastrophe produce an alteration of the natural evolutionary sequence for clusters with w0 ≥ 6.9. from the comparison with extragalactic gcs systems we have deduced that, in the case of lmc, smc and fornax system, these clusters have a gaussian like distribution around a peak value w0 ∼ 5. mw and m31 system are very similar in their features and w0 distribution, with low-w0 tail and a peak value in corrispondence of the stability limit. we can assume this as the main product of disk shocking, as well as extragalactic capture and low speed evolution objects mentioned above. for ngc5128 there is no low w0 region tail, but only a narrow peaked distribution around the stability limit. references [1] aguilar, l., hut, p., ostriker, j.p.: 1988, apj, 335, 720 doi:10.1086/166961 [2] barmby, p., mclaughlin, d.e., harris, w.e., gretchen, l.h., forbes, d.a.: 2007, aj, 133, 2764 [3] bica, e., bonatto, c., barbuy, b., ortolani, s.: 2006, aap, 450, 105 doi:10.1051/0004-6361:20054351 [4] cohn, h., hut, p.: 1984, apj, 277, l45 doi:10.1086/184199 [5] dinescu, d.i., terrence, m.g., van altena, w.f.: 1999, aj, 117, 1792 [6] dotter, a. et al.: 2010, aj, 708, 698 [7] gnedin, o.y., ostriker, j.p.: 1997, aj, 474, 223 [8] gòmez, m., geisler, d., harris, w.e., richtler, t., harris, g.l.h., woodley, k.a.: 2006, aap, 447, 877 doi:10.1051/0004-6361:20053393 [9] harris, w.e.: 1996, aj, 112, 1487 [10] hut, p., djorgovski, s.: 1992, nat, 359, 806 doi:10.1038/359806a0 [11] katz, j.: 1980, mnras, 190, 497 doi:10.1093/mnras/190.3.497 [12] king, i.: 1966, aj, 71, 64 [13] mackey, a.d., gilmore, g.f.: 2003a, mnras, 338, 85 doi:10.1046/j.1365-8711.2003.06021.x 238 http://dx.doi.org/10.1086/166961 http://dx.doi.org/10.1051/0004-6361:20054351 http://dx.doi.org/10.1086/184199 http://dx.doi.org/10.1051/0004-6361:20053393 http://dx.doi.org/10.1038/359806a0 http://dx.doi.org/10.1093/mnras/190.3.497 http://dx.doi.org/10.1046/j.1365-8711.2003.06021.x data analysis of globular cluster harris catalogue... part ii [14] mackey, a.d., gilmore, g.f.: 2003b, mnras, 338, 120 doi:10.1046/j.1365-8711.2003.06022.x [15] mackey, a.d., gilmore, g.f.: 2003c, mnras, 340, 175 doi:10.1046/j.1365-8711.2003.06275.x [16] mackey, a.d., gilmore, g.f.: 2004, mnras, 355, 504 doi:10.1111/j.1365-2966.2004.08343.x [17] quinlan, g.d.: 1996, newa, 1, 255 [18] van den bergh, s.: 2011, pasp, 123, 1044 doi:10.1086/662132 [19] zinn, r.: 1985, apj, 293, 424 239 http://dx.doi.org/10.1046/j.1365-8711.2003.06022.x http://dx.doi.org/10.1046/j.1365-8711.2003.06275.x http://dx.doi.org/10.1111/j.1365-2966.2004.08343.x http://dx.doi.org/10.1086/662132 introduction discussion conclusions 311 acta polytechnica ctu proceedings 1(1): 311–315, 2014 311 doi: 10.14311/app.2014.01.0311 sifap: a new fast astronomical photometer filippo ambrosino1, franco meddi1, roberto nesci 2, corinne rossi1, silvia sclavi1, ivan bruni3 1dipartimento di fisica università la sapienza, p.le a. moro 5, 00185 roma, italy, 2inaf-iaps, tor vergata, roma, italy, 3istituto nazionale di astrofisica, osservatorio astronomico, via ranzani 1, 40127 bologna, italy. corresponding author: filippo.ambrosino@roma1.infn.it abstract a fast photometer based on sipm technology was developed and tested at the university of rome “la sapienza” and at the bologna observatory. in this paper we present the improvements applied to our instrument, concerning new cooled sensors, a new version of the electronics and an upgraded control timing software. keywords: instrumentation: photometer stars: pulsars. 1 instrument description the general description of our 3-channel photometer sifap (silicon fast astronomical photometer) is presented in meddi et al. (2011; 2012). we only remind here that each channel is dedicated to a different target: variable source (channel 0), nearby sky (channel 1) and reference star (channel 2). a gps receiver sends a pulse per second (pps) to a micro processor (µp) which in turn produces a signal that we call gated pps. this last is used both to drive two leds in order to have an optical temporal marker and to synchronize our custom electronics. during the last year we replaced the old hamamatsu multi pixel photon counter (mppc) modules with new ones, having the sensor cooled by a built-in peltier cell 1; we modified our electronics and to increase the uninterrupted acquisition time and to reduce the sampling gate duration from 0.55 ms down to 0.1 ms. the new block diagram of the instrument is shown in figure 1. the thermo-electric cooled system allows to reach a fixed working temperature of -10 ◦c, with a large reduction of the mean dark count. the s/n ratio is improved by a factor 3 with respect to the previous one. the modules have limited geometrical dimensions and low weight so they can be directly located at the exit pupil of the telescope. other characteristics of the new sensors are similar to the old ones: pixel size of 50 µm for a total active area of 1 mm2 and photon detection efficiency with a maximum value of about 50% at wavelength of 440 nm2. the built-in electronics of each hamamatsu module can generate three types of output: pulse count via usb interface already processed in time windows of 1 ms, analog and discriminated. we use the last one to feed our electronics which uses a new protocol for data exchange between a field programmable gate array and a µp and a new data storage support. the integration time windows for this output are now 0.1 ms. we called this electronics p3e, which stands for pulsar pulse period extractor. the pointing and the signal maximization procedures and the fast pre-analysis to check the pulsating behavior of the pointed source are described in meddi et al. (2012). 1 http://www.hamamatsu.com/resources/pdf/ssd/c11208 series kacc1176e03.pdf 2http://www.hamamatsu.com/resources/pdf/ssd/s11028 series kapd1026e04.pdf 311 http://dx.doi.org/10.14311/app.2014.01.0311 filippo ambrosino et al. figure 1: general block diagram of sifap mounted at the telescope. the gps antenna is located outside the dome. hamamatsu mppc sensors are integrated in their modules, p3e units are fed by the discriminated output. each couple of mppc sensor and p3e unit is dedicated to a different target: variable source (mppc0 and p3e0), nearby sky (mppc1 and p3e1) and a reference star (mppc2 and p3e2). figure 2: fft applied to the whole raw data acquired by p3e0 on 2012, december 19. 312 sifap: a new fast astronomical photometer figure 3: fundamental frequency and period values of the crab pulsar signal obtained by our data with applied the heliocentric corrections compared with those computed from jbo ephemeris. the uncertainties on the jbo values are due to a numerical interpolation procedure. the quoted errors on our data are statistical only. figure 4: crab pulsar light curve folded by the xronos task efold for p3e0 (2012, december 19) corrected data. 2 data analysis and results on 2012 december 19 we observed the crab pulsar with the loiano telescope, for one hour. to optimize the pointing of the target, we performed a quick pre-processing analysis on a short acquisition of the data (∼ 100 s), based on autocorrelation and fft techniques. we then analyzed the whole data set with a fft analysis to estimate the spin period of the pulsar. in figure 2 we show the amplitude spectrum of the crab pulsar signal obtained from data collected 313 filippo ambrosino et al. by p3e0. we used the xronos package of the high energy astrophysics science archive research center (heasarc) to apply the heliocentric correction (task earth2sun) and to determine the best fitting period for the crab pulsar signal (task efsearch). in figure 3 we compare the jodrell bank observatory (jbo) ephemeris3 prediction for the fundamental frequency and period of the crab pulsar signal with the estimates of the same quantities obtained by our barycentered data. the agreement between the two results is within 300 µhz for the fundamental frequency and 300 ns for the spin period. such discrepancy can not be justified without taking into account systematic effects, which are explained in terms of the clock drift which depends on the temperature. this drift is estimated by measuring the time interval between the two optical markers generated by the gated pps signal, the first at the beginning and the second one at the end of the acquisition. taking into account such effect, the differences for the spin period are reduced below 15 ns. systematic uncertainties are due to i) the propagation of the gated pps signal trough the cables and ii) the rising edge of leds. such systematic uncertainties produce a total delay of about 320 ns. finally, the accuracy on the pulse output time of the pps signal which is ±0.001 ms at the rising edge of the pulse itself. all these effects are not included in the numerical computation of the uncertainty on the period obtained by applying the barycentric correction. the crab light curves were obtained by using the task efold on the barycentered data; the result for p3e0 is shown in figure 4. the shape of the primary and the second peak, and their flux ratio, are in good agreement with the literature data (e.g. golden et al., 2000; zampieri et al., 2011). the sky and reference star measurements were processed in the same way as the ones belonging to the target to look for spurious signals which might interfere with the astronomical signal from the crab nebula: no evidence of periodicity was detected. 3 conclusions we built a fast photometer able to collect data of periodic signals with high time accuracy integrating in time windows down to 0.1 ms. with the loiano 1.5 m telescope we derived the period and a high s/n light curve of the crab pulsar with a measurement duration of about 1 hour. in these conditions we obtained a fair agreement with the fundamental frequency and the spin period calculated from the jbo database. we intend to upgrade our custom electronics aiming at reaching shorter time sampling keeping a good s/n ratio. our final goal would be to compare fast optical measurements with the γ-ray, x-ray, ir and radio ones to explore more deeply the structure and the phenomena occurring in the emitting regions of pulsars. in particular, we want to study how the spin period slows down (dissipative process) and the amount of the eventual phase delay among the peaks in the various band (phase shifting process). we need to collect high quality data with a 0.001 ms time windows to be able to perform these kind of investigations. to this purpose we modified our gps system to reach a more accurate determination of the absolute time scale (within 100 ns) by using a burst of n gated pps signals instead of the single one presently used. we started with theoretical computation of the minimum possible time sampling for a v ∼ 16 mag object as a function of the telescope diameter. the results indicate that it could be possible to measure objects of the same magnitude as the crab pulsar with similar s/n ratio in time windows shorter than 0.1 ms. for instance, with a 5 m telescope the integration time could be reduced down to 0.01 ms. moreover it is mandatory to have larger telescopes in order to have enough detectable photons and higher resolution on the absolute timing by adopting a military class gps antenna which would be able to reach at least 1 ns accuracy. in this case it would be necessary to upgrade the optical time marker by substituting standard leds with laser ones. acknowledgement the authors thank the bologna observatory for the logistic support and the technical assistance during the observations. the universita’ di roma “la sapienza” supported the project. this research has made use of the simbad database operated at cds, strasbourg, france. references [1] golden a., shearer a., redfern r. m., beskin g. m., neizvestny s. i. et al., 2000, a&a, 363, 617 [2] meddi f., ambrosino f., rossi c., nesci r., sclavi s. et al., 2011, acta polytechnica, 51, n.6, 42. [3] meddi f., ambrosino f., rossi c., nesci r., sclavi s. et al., 2012, pasp, 124, 448. doi:10.1086/665925 [4] zampieri l., germana c., barbieri c., naletto g., cade a. et al., 2011, adv. space res., 47, 365. doi:10.1016/j.asr.2010.07.016 3http://www.jb.man.ac.uk/pulsar/crab.html 314 http://dx.doi.org/10.1086/665925 http://dx.doi.org/10.1016/j.asr.2010.07.016 sifap: a new fast astronomical photometer discussion b. aschenbach: how do you explain that the second harmonic in the amplitude spectrum is higher than the first one? f. ambrosino: the amplitude of the single harmonic depends on the shape of the light curve; in the case of the crab pulsar in one period there are two emission peaks of different intensity and with a phase shift of about 0.45. 315 instrument description data analysis and results conclusions 242 acta polytechnica ctu proceedings 2(1): 242–245, 2015 242 doi: 10.14311/app.2015.02.0242 line evolution of the nova v5587 sgr from early to nebula phase t. kajikawa1, a. arai1,2, m. nagashima1, h. kawakita1, m. yamanaka3,4,5, k. kawabata3, s. kiyota6 1kyoto sangyo university, koyama astronomical observatory 2university of hyogo, nishi-harima astronomical observatory 3hiroshima university 4kyoto university 5konan university 6vsolj corresponding author: kajikawat8@gmail.com abstract the spectral evolution of the nova v5587 sgr has been monitored at koyama astronomical observatory and higashihiroshima observatory, japan, from the early to nebula phase. the nova rebrightened several times. the spectra during the early phase showed emission lines of hα, hβ, o i, he i, he ii, n ii, fe ii. nova v5587 sgr is classified into the fe ii type. the helium abundance of the nova is estimated as n(he)/n(h) = 0.134 ± 0.09. the light curve, the spectral evolution, and the helium abundance in v5587 sgr are similar to those of the nova pw vul. keywords: cataclysmic variables classical novae optical spectroscopy photometry individual: v5587 sgr. 1 introduction photometric observations of novae are usually performed by many observers worldwide. the light curves of the novae have been revealed in detail, and they can be classified into several classes (strope et al. 2010). in contrast, spectroscopic observations of novae are not performed as frequently as photometric observations because (i) the number of available spectrographs (and observers) are limited; (ii) some fraction of novae are too faint to observe by spectrographs in their later phase (or even at the maximum for some novae). as novae show a variety of spectral evolutions, spectroscopic monitoring can provide new insights into the physics and chemistry of novae. for example, recently nagashima et al. (2013) detected molecular absorption bands of c2 and cn in nova v2676 oph. (this is the first detection of c2 in novae and the second for cn.) this detection was performed as a part of long-term spectroscopic observations of this nova. the light curves of novae in some classes (strope et al. 2010) can be successfully reproduced by models (as reviewed by prof. kato at this meeting). however, the light curves in the “jitter”and “oscillation”classes (strope et al. 2010) could not be reproduced well by the models. in this paper, we focus on the “jitter”class and introduce our recent spectroscopic observations of v5587 sgr classified in this class. we discuss the nature of this nova as it is likely that there are differrent types of oscillations (i.e., different physical conditions) of novae in the “jitter”class. different types of novae might 6 8 10 12 14 16 18 0 100 200 300 400 500 v ( m a g n itu d e s) days after discovery (t) v5587 sgr pw vul figure 1: v band magnitudes of v5587 sgr and pw vul. the light curve of v5587 sgr is from higashihiroshima observatory, vsnet, and aavso observations. the light curve of pw vul was refer to the online data of strope et al.(2010). 242 http://dx.doi.org/10.14311/app.2015.02.0242 line evolution of the nova v5587 sgr from early to nebula phase be mixed together in this class. even though there are high-quality photometric and spectroscopic observations for some jitter class novae, the mechanism for their rebrightening is still unclear (tanaka et al. 2011). by frequent spectroscopic observations, we can observe phenomena useful to reveal the physics of the nova. in order to investigate the mechanism for the rebrightening in the “jitter”class novae, we have conducted spectroscopic observations of v5587 sgr in 2011 – 2012. the nova v5587 sgr (= nova sgr 2011 no.1) was discovered on ut 2011 january 25.86 by nishimura (nakano et al. 2011). the first low-dispersion spectrum was obtained on ut 2011 january 28 at koyama astronomical observatory. the spectrum shows prominent emission lines of hα, hβ and o i. these features suggest that this object is a classical nova (arai 2011; imamura 2011). 2 observations our photometric and spectroscopic observations were performed at two sites. one site is koyama astronomical observatory (kyoto sangyo university, kyoto, japan). we used the 1.3m araki telescope with the losa/f2 spectrograph (arai et al., in prep.) for low-dispersion spectroscopy. the wavelength coverage is 400 – 800 nm and the spectral resolving power is r = λ/∆λ ∼ 580 at 600 nm. the other site is higashi-hiroshima observatory (hiroshima university, hiroshima, japan). we used the 1.5m kanata telescope with trispec (triple range imager and spectrograph with polarimeter, shutdown in 2012) for photometry and with howpol (hiroshima one-shot wide-field polarimeter) for spectroscopy. its wavelength coverrage is 400 – 1000 nm, and spectral resolving power is r ∼ 400 at 600 nm. 450 500 550 600 650 700 750 800 850 n o rm a liz e d f lu x + c o n st . wavelength (nm) +596d +229d +213d +185d +167d +154d +147d +129d +100d +93d +81d +69d +68d +61d +60d +45d +43d +39d +35d +27d +11d +9d +8d h e i i (4 6 8 .6 ) h β f e i i( 4 2 ) (4 9 2 .4 ) f e i i( 4 2 ) (5 0 1 .8 ) f e i i( 4 2 ) (5 1 6 .9 ) [o i ii ] (4 9 5 .9 ) [o i ii ] (5 0 0 .6 ) n i i (5 6 8 .0 ) [n i i] ( 5 7 5 .5 ) h e i ( 5 8 7 .4 ) [o i ] (6 3 0 .0 ) [o i ] (6 3 6 .3 ) h α h e i ( 6 6 8 .1 ) h e i ( 7 0 6 .5 ) [o i i] ( 7 3 3 .0 ) o i ( 7 7 7 .4 ) figure 2: the normalized spectra of v5587 sgr observed from +8d (= ut 2011 february 2) to +596d (= ut 2012 september 12) after the discovery. these spectra were obtained at koyama astronomical observatory except for the data on +35d and +45d (= ut 2001 march 3 and 12), taken at higashi-hiroshima observatory. 243 t. kajikawa et al. 3 results figure 1 shows light curves of v5587 sgr. since v5587 sgr showed an erratic variation in the light curve, we consider that the nova is classified in the “jitter”class. particularly, v5587 sgr seems to be a member of the sub-class prototyped by pw vul as shown in strope et al. (2010). based on the light curves, we determined the date of the visual maximum as 9 days after discovery and also determined t2 = 12 days and t3 = 108 days. from (b v) at maximum magnitude and at t2, we estimate the color excess by intersterllar extinction; e(b v) = 0.85 – 1.26. the light curve of v5587 sgr is quite similar to that of the “jitter”class nova pw vul as shown figure 1 (schwarz et al. 1997, and references therein). pw vul was discovered on 1984 july 27 (= jd 2445917) by wakuda (kosai 1984). this nova reached its maximum light of 6.3 magnitude in the v band on 1984 aug 4th. we compare v5587 sgr with pw vul from various viewpoints in the next section. 500 550 n o rm a liz e d f lu x wavelength (nm) +45d +43d +39d +35d +27d +11d +9d +8d hβ fe ii (42) figure 3: the spectra of v5587 sgr in early phase. the nova never showed any regrowth of a p-cygni profile. figure 2 shows the growth of hα, hβ, he i, he ii, n ii, [n ii], o i, [o i], [o ii], and fe ii emission lines of v5587 sgr. novae that exhibit rebrightening in the early phase often show the regrowth of a p-cygni profile (tanaka et al. 2011, csak et al. 2005) as in v4745 sgr. however, v5587 sgr has never shown this behavior (figure 3). it is considered that the nova entered into the nebula phase between july 11 to july 29 because the [o iii] forbidden emission lines dominated hβ at that time (figure 4). we assume that the electron density in the nebula phase was ne = 10 6 ∼ 108 cm−3 according to iijima & esenoglu (2003) and iijima (2006), and we estimate the helium abundance of the ejecta based on the data between july 11 and july 29 ; n(he)/n(h) = 0.134 ± 0.09 (for ne = 106 cm−3), n(he)/n(h) = 0.139 ± 0.09 (for ne = 108 cm−3). 450 500 550 600 650 700 750 800 n o rm a liz e d f lu x wavelength (nm) +147d +129d h e i i (4 6 8 .6 ) h β [o i ii ] + f e i i( 4 2 ) n i i (5 6 8 .0 ) h e i ( 5 8 7 .4 ) [o i ] (6 3 0 .0 ) [o i ] (6 3 6 .3 ) h e i ( 6 6 8 .1 ) h e i ( 7 0 6 .5 ) [o i i] ( 7 3 3 .0 ) o i ( 7 7 7 .4 ) figure 4: the spectra of v5587 sgr in nebular phase(+129 days and +147 days after its maximum). an emission line of he ii was recognized in these spectra. 4 discussion & conclusion the decline rates, amplitudes, and intervals of rebrightening observed in v5587 sgr are similar to those of pw vul. this fact suggests that the wd mass and physical parameters related to the explosion of v5587 sgr would be similar to those of pw vul. comparison between the spectra of v5587 sgr and pw vul also lead us to the same conclusion, namely, they are similar to each other, e.g., the late-phase spectra of v5587 sgr (shown in figure 4 in rosino and iijima 1987). figure 5 also shows helium abundances vs t3 in 20 various novae. helium abundance of v5587 sgr seems to be similar to that of pw vul. this figure indicates that helium abundance is not correlated with t3. in summary, we have performed photometric and low-dispersion spectroscopic observations of v5587 sgr from early to nebula phase. the photometry showed erratic variations of the light curve. the spectra during the early phase showed emission lines of hα, hβ, and fe ii (i.e., fe ii type) and v5587 sgr never showed regrowth of a p-cygni profile. the nova entered the nebula phase between july 11 to july 29 (t = 158 to 178days). we estimated that the helium abundance of v5587 sgr is n(he)/n(h) = 0.134 ± 0.09. the nova is very similar to pw vul considering decline rates, spectral features, and helium abundance. we obtained a new sample of the “jitter”class novae. 244 line evolution of the nova v5587 sgr from early to nebula phase -1 -0.5 0 0.5 0.5 1 1.5 2 2.5 3 3.5 lo g 1 0 (n (h e )/ n (h )) log10t3 v1494 aql(1) v723 cas(2) v2468 cyg(3) cl aql(4) v407 cyg(5) v1500 cyg(6) qu vul(7) v351 pup(8) u sco(9) pw vul(10),(11) v842 cen(11) v827 her(11) qv vul(11) v2214 oph(11) v977 sco(11) v443 sct(11) v1668 cyg(11) v693 cra(11) v1370 aql(11) v5587 sgr figure 5: the helium abundance of 20 novae. references: (1) iijima & esenoglu (2003); (2) iijima (2006); (3) iijima & naito (2011); (4) iijima (2012a); (5) iijima (2012b); (6) ferland (1978); (7) schwarz (2002); (8) saizar (1996); (9) iijima (2002); (10) schwarz et al. (1997); (11) andreä et al. (1994). references [1] andreä, j., drechsel, h., starrfield, s.: 1994, a&a 291, 869. [2] arai, a.: 2011, iauc 9196, 2. [3] chochol, d., et al.: 2012, mmsai 83, 767. [4] ferland, g.j.: 1978, apj 219, 589. doi:10.1086/155818 [5] iijima, t.: 2002, a&a 387,1013. [6] iijima, t., esenoglu, h.h.: 2003, a&a 404, 997. [7] iijima, t.: 2006, a&a 451, 563. [8] iijima, t., naito, h.: 2011, a&a 526a, 73. [9] iijima, t.: 2012a, mmsai 83, 835. [10] iijima, t.: 2012b, a&a 544a, 26. [11] imamura, k.: 2011, iauc 9196, 2. [12] nagashima, m., et al.: 2014, journal?, in press. [13] nakano, s., nishimura, h., kiyota, s., yusa, t.: 2011, iauc 9196, 1. [14] rosino, l., iijima, t.: 1987, ap&ss 130, 157. doi:10.1007/bf00654989 [15] saizar, p., et al.: 1991, apj 367,310. doi:10.1086/169630 [16] saizar, p., et al.: 1996, mnras 279, 280. doi:10.1093/mnras/279.1.280 [17] schwarz, g.j., et al.: 1997, mnras 290, 75. doi:10.1093/mnras/290.1.75 [18] schwarz, g.j.: 2002, apj 577, 940. doi:10.1086/342234 [19] strope, r.j., schaefer, b.e., henden, a.a.: 2010, aj 140, 34. [20] tanaka, j., et al.: 2011, pasj 63, 159. [21] tanaka, j., et al.: 2011, pasj 63, 911. 245 http://dx.doi.org/10.1086/155818 http://dx.doi.org/10.1007/bf00654989 http://dx.doi.org/10.1086/169630 http://dx.doi.org/10.1093/mnras/279.1.280 http://dx.doi.org/10.1093/mnras/290.1.75 http://dx.doi.org/10.1086/342234 introduction observations results discussion & conclusion 123 acta polytechnica ctu proceedings 1(1): 123–126, 2014 123 doi: 10.14311/app.2014.01.0123 a multi-wavelength view of the xmm-newton galactic plane ada nebot gómez-morán1, christian motch1, on behalf of the xmm-newton survey science centre 1observatoire astronomique de strasbourg, université de strasbourg, cnrs, umr 7550, 11 rue de l’université, 67000 strasbourg, france. corresponding author: ada.nebot@astro.unistra.fr abstract we present an x-ray survey of the galactic plane conducted by the survey science centre of the xmm-newton satellite. the survey contains more than 1300 x-ray detections at low and intermediate galactic latitudes and covering 4 deg2 well spread in galactic longitude. from a multi-wavelength analysis, using optical spectra and helped by optical and infrared photometry we identify and classify about a fourth of the sources. the observed surface density of soft x-ray (< 2 kev) sources decreases with galactic latitude and although compatible with model predictions at first glance, presents an excess of stars, likely due to giants in binary systems. in the hard band (> 2 kev) the surface density of sources presents an excess with respect to the expected extragalactic contribution. this excess highly concentrates towards the direction of the galactic centre and is compatible with previous results from chandra observations around the galactic centre. the nature of these sources is still unknown. keywords: stars binaries spectroscopy photometry ir optical x-rays. 1 introduction galactic x-ray surveys can help us to learn about the galaxy’s structure, stellar formation and evolution. since long we know that the soft x-ray emission (< 2 kev) of the galaxy is dominated by stars. but the x-ray luminosity function of hard x-ray emitting sources (>2 kev) is still not well constrained, in particular at low to intermediate x-ray luminosities (10 < lx < 10 34erg s−1). cataclysmic variables and active binaries contribute at x-ray luminosities up to ∼ 1033 erg s−1(sazonov et al. 2006), but recent results show that other sources, such as massive ob stars, wolf-rayet and wind colliding binaries also contribute to this lx range (mauerhann et al. 2010, anderson et al. 2011). a census of these populations is thus needed. discrepancies between observations and model predictions exist since the first x-ray surveys of the galaxy were carried out by the einstein satellite. stellar population models predict less stars than we observe, in particular, there is an excess of stars with spectral type k, this is the so called “yellow excess”. the origin of this excess is highly debated and there are two possible explanations: either these k stars are young stars or old evolved giants, which can be either single or in binary systems (favata et al. 1988, sciortino et al. 1995). x-ray surveys can reveal concentrations of specific populations in certain regions of the galaxy: high-mass x-ray binaries (hmxb) seem to be concentrated in the norma arm (lutovinov et al. 2005); the gould belt contains a young population of late type stars (guillout et al. 1998). last but not least, different populations of theoretically predicted x-ray binaries are still to be discovered. binary population models predict a population of x-ray binaries at low luminosities, prehmxbs and pre-low-mass x-ray binaries that would be accreting matter coming from the wind of the companion star at low rates rather than from roche lobe overflow (willems & kolb 2003, pfahl et al 2002). other predicted objects are hmxbs with white dwarf as compact object instead of a classical hmxb where the compact object is a neutron star (motch et al 2007). so far, none of these objects is confirmed observationally. to learn about all these questions we need to study the x-ray content of the galaxy. 2 xmm-newton survey science centre the xmm-newton satellite can help us to give answer to these questions. thanks to it’s large field of view and it’s high sensitivity, each xmm-newton pointed observation discovers from 30 to 100 serendipitous sources. the xmm-newton survey science centre (ssc) generates catalogues of serendipitous sources and publishes them on a regular basis. the latest available catalogue, the 2xmmi-dr3 catalogue, contains about 250 000 unique sources, and the 3xmm catalogue will 123 http://dx.doi.org/10.14311/app.2014.01.0123 ada nebot gómez-morán, christian motch soon be released containing more than half a million detections corresponding to about 370 000 unique sources. one of the goals of the ssc is to characterise and identify these sources in an statistical manner. for that, specific surveys have been studied in great detail, such as a bright flux sample by della ceca et al. 2004, a medium flux sample by barcons et al. 2002 and the galactic survey from hands et al. 2004, with spectroscopic identification carried out by motch et al 2010. these surveys serve as learning samples for an statistical identification. in this paper we briefly describe the main results of a survey carried out by the xmmnewton ssc in the galactic plane. for details on the survey we refer the reader to nebot gómez-morán et al. 2013. epic-pn wr110 dss 2mass figure 1: example field studied in this survey. from left to right, x-ray (epic-pn), optical (dss) and infrared (2mass) images around the wr 110. 3 the galactic plane survey the xmm-newton ssc survey of the galactic plane contains 26 pointings, covering about 4◦ at low and intermediate galactic latitudes (|b| < 20◦) and well spread in galactic longitude. it contains about 1300 detections up to limiting fluxes of ∼ 2 × 10−15 erg cm−2s−1in the soft (kt < 2 kev) band and ∼ 10−14 erg cm−2s−1in the hard band (2 < kt < 12 kev). 3.1 multi-wavelength identification in order to identify these sources in the optical and the infrared we cross-correlated the x-ray positions with large optical and infrared published catalogues, such as the sdss, usno and 2mass. we show an example of one of the fields studied in this survey in figure 1. the xmm-newton detected about 50 sources around the pointed observation of wr 110. if we look in the optical (middle panel) the field clearly contains a large number of sources, while in the infrared (right panel) the field becomes very crowded. this example reflects the difficulty of multi-wavelength identification of x-ray sources in the galactic plane. due to the relatively large x-ray positional errors, and the high surface density of sources at faint fluxes, only not too distant and relatively bright sources can be identified in the optical and the infrared. among the ∼1300 x-ray detected sources we could find counterparts in the optical to about 60%, number which drops to about 40% in the infrared. 3.2 source classification a spectroscopic campaign was carried out to classify the sources. about one forth of the detected x-ray sources are classified. most of them are active coronae, but we also found other interesting objects, such as young t tauri and herbig-ae stars, and old cataclysmic variables. we identify four be stars with x-ray luminosities around 1032 erg s−1 a value too high for normal be stars. these systems thus belong to the class of γ-cas analogues, which means that we have almost doubled the number of γ-cas objects known until the date (see talk by c. motch in this congress). we cross-matched our x-ray sources with other published catalogues making use of vizier and the xcat-db databases, and produced spectral energy distributions which helped us to discern the type of source. infrared data from the wise catalogue revealed one of the t tauri has a transition disk, which highlights the relevance of having a multi-wavelength view of high energy sources. 124 a multi-wavelength view of the xmm-newton galactic plane ‣ ‣ ‣ 10 -15 10 -14 10 -13 10 -12 s (0.5-2 kev) [erg cm -2 s -1 ] 10 -1 10 0 10 1 10 2 n (> s ) [s o u rc e s /d e g 2 ] |b| = 0 o |b| = 15 o |b| = 50 o |b| = 60 o b = 50º b = 60º b = 15º b = 0º 0.5-2.0 kev observed figure 2: logn – logs curves in the soft band (< 2 kev) and as a function of galactic latitude. we included the data from lopez-santiago et al. 2007 and barcons et al. 2007 at high galactic latitudes. using infrared colour-colour diagrams for the stellar content of our survey we could distinguish main sequence stars from evolved giants. on the one hand, due to magnetic braking, stellar rotation decreases with the age of the stars. on the other hand the x-ray luminosity is strongly related to rotation. this implies that normal old giant stars are not expected to be strong x-ray emitters. a possible explanation for the origin of the x-ray emission of these evolved stars could be if they are in synchronised binary systems, where the rotation period is equal to the orbital period, breaking thus the relation between x-ray emission and age. to test this hypothesis we calculated the x-ray colours expected for stars of different age and for binaries, in particular for rs cvn and by draconis. we found that observed xray and infrared colours for dwarfs are compatible with young to intermediate age stars, while giant stars have colours compatible with rs cvn binaries, i.e. binaries where at least one of the stars is a giant star. the giant stars we find in our survey mostly have k spectral type, i.e. they are “yellow stars”. 4 results we computed the logn – logs curves in the soft band, i.e. the surface density of sources as a function of the sensitivity of our survey, for different bins of galactic latitude. we find that the number of stars per square degree decreases with galactic latitude. although this result is not a surprise, it’s the first time that is shown in x-rays. we also see that the x-ray luminosity function has a varying slope with galactic latitude. we compared our results with the expected curves from a modified version of the galactic x-ray model from guillout et al. 1996. at first glance we obtained compatible results, a varying slope of the x-ray luminosity function and a decreasing number of sources towards higher galactic latitudes, reflecting the different scale-height of stars and the relative contribution of different populations. but for a given flux, the number of observed stars is higher than the predicted value. this result is not surprising since binaries have not been taken into account in our model. in other words, the excess of observed sources with respect to model predictions is due to the yellow giant stars that we have found in our survey, and are likely in binary systems. 10 -14 10 -13 10 -12 10 -11 s (2-12 kev) [erg cm -2 s -1 ] 10 -1 10 0 10 1 10 2 10 3 n (> s ) [s o u rc e s /d e g 2 ] galactic center motch et al 2010 10 32 10 33 10 34 luminosity (2-12 kev) [erg s -1 ] 2-12 kev gc region (l,b) ~ (1º,0º) ‣ ‣ ‣ hands+2004, motch+2010 (l,b) =(20º,0º) b ~ 0º figure 3: logn – logs curves of hard sources (> 2 kev) in the direction of the galactic center. the expected extragalactic contribution from mateos et al. 2008 has been subtracted. we carried out the same exercise for hard sources, we constructed logn – logs curves for sources harder than 2 kev. since stars are not expected to contribute at this high energies, we subtracted the expected extragalactic contribution. the logn – logs curves present an excess of sources with respect to the expected extragalactic contribution close to the galactic center region studied in this survey (l ∼ 0.9◦, b ∼ 0◦). we compared this result with that obtained by motch et al. 2010 at l ∼ 20◦, b ∼ 0◦ and obtained that the number of sources per square degree observed at a given flux decreases steeply with galactic longitude in the direction of the galactic center. we obtained consistent results with hong et al. 2009 based on chandra observations of seven fields around the galactic center. this sources contribute to the galactic ridge emission, initially thought to be of diffuse origin. the nature of these sources is unknown, but if this population would be associated to the galactic center itself, the x-ray luminosities obtained are in the range 1033−1034 erg s−1, i.e. values which are comparable with pre-hmxbs and pre-lmxbs. 125 ada nebot gómez-morán, christian motch acknowledgement we would like to thank the organisers of this conference for their work and for giving us the opportunity to show our results. references [1] anderson l. d., bania t. m., jackson j. m., clemens d. p., heyer m., simon r., shah r. y., rathborne j. m., 2009, apjs, 181, 255 doi:10.1088/0067-0049/181/1/255 [2] barcons, x., carrera, f. j., watson, m. g., et al. 2002, a&a, 382, 522 [3] barcons, x., carrera, f. j., ceballos, m. t., et al. 2007, a&a, 476, 1191 [4] della ceca, r., maccacaro, t., caccianiga, a., et al. 2004, a&a, 428, 383 [5] favata, f., sciortino, s., rosner, r., & vaiana, g. s. 1988, apj, 324, 1010 doi:10.1086/165957 [6] guillout, p., haywood, m., motch, c., & robin, a. c. 1996, a&a, 316, 89 [7] guillout, p., sterzik, m. f., schmitt, j. h. m. m., motch, c., & neuhaeuser, r. 1998, a&a, 337, 113 [8] hands, a. d. p., warwick, r. s., watson, m. g., & helfand, d. j. 2004, mnras, 351, 31 doi:10.1111/j.1365-2966.2004.07777.x [9] hong, j. s., van den berg, m., grindlay, j. e., & laycock, s. 2009, apj, 706, 223 doi:10.1088/0004-637x/706/1/223 [10] lópez-santiago, j., micela, g., sciortino, s., et al. 2007, a&a, 463, 165 [11] lutovinov a., revnivtsev m., gilfanov m., shtykovskiy p., molkov s., sunyaev r., 2005, a&a, 444, 821 [12] mauerhan j. c., muno m. p., morris m. r., stolovy s. r., cotera a., 2010, apj, 710, 706 [13] mateos, s., warwick, r. s., carrera, f. j., et al. 2008, a&a, 492, 51 [14] motch, c., lopes de oliveira, r., negueruela, i., haberl, f., & janot-pacheco, e. 2007, in astronomical society of the pacific conference series, vol. 361, active ob-stars: laboratories for stellare and circumstellar physics, ed. a. t. okazaki, s. p. owocki, & s. stefl, 117 [15] motch, c., warwick, r., cropper, m. s., et al. 2010, a&a, 523, a92 [16] nebot gómez-morán a. et al., 2013, a&a, 553, a12 [17] pfahl, e., rappaport, s., & podsiadlowski, p. 2002, apj, 571, l37 doi:10.1086/341197 [18] sazonov, s., revnivtsev, m., gilfanov, m., churazov, e., & sunyaev, r. 2006, a&a, 450, 117 [19] sciortino, s., favata, f., & micela, g. 1995, a&a, 296, 370 [20] willems, b. & kolb, u. 2003, mnras, 343, 949 discussion david buckley: can you give the percentages of the different optical counterpart types of the ∼ 300 sources identified? ada nebot: among the 316 classified sources there are three t tauri stars, one herbig ae star, two cataclysmic variables, and four γ-cas analogues. a few sources are extragalactic (in the highest galactic latitude fields), and the remaining sources are stars, where about 10% are giants. classification of giants was only done for stars with the best 2mass photometry, meaning that the number of giants in our survey could be higher. 126 http://dx.doi.org/10.1088/0067-0049/181/1/255 http://dx.doi.org/10.1086/165957 http://dx.doi.org/10.1111/j.1365-2966.2004.07777.x http://dx.doi.org/10.1088/0004-637x/706/1/223 http://dx.doi.org/10.1086/341197 introduction xmm-newton survey science centre the galactic plane survey multi-wavelength identification source classification results 264 acta polytechnica ctu proceedings 2(1): 264–268, 2015 264 doi: 10.14311/app.2015.02.0264 on the 2011 outburst of the recurrent nova t pyxidis l. izzo1, m. della valle2, a. ederoclite3, m. henze4 1dip. di fisica, sapienza university of rome, p.le a. moro 2, rome; icranet, p.le della repubblica 10, pescara, italy 2inaf osservatorio di capodimonte, salita moiariello 16, 80131 napoli, italy 3centro de estudios de fisica del cosmos de aragón, plaza san juan 1, planta 2, teruel, e44001, spain 4european space astronomy centre, p.o. box 78, 28692 villanueva de la cañada, madrid, spain corresponding author: luca.izzo@gmail.com abstract we discuss the nebular phase emission during the 2011 outburst of the recurrent nova t pyxidis and present preliminary results on the analysis of the line profiles. we also present some discussions about the binary system configurations and the x-ray emission, showing that the white dwarf mass should be larger than 0.8 m�. keywords: spectroscopy x-rays individual: t pyx. 1 introduction the 2011 outburst of the recurrent nova t pyx shed more light on some interesting characteristics of this peculiar nova. the optical evolution was not very different from the last (1966) outburst, showing the 8 days ’shoulder’ before the final rise to the maximum luminosity, which was reached ∼ 30 days from the initial outburst. the distance of t pyx was determined from light echoes to be 4.8 ± 0.5 kpc [1], which is marginally consistent within the errors with the value of 3.5 ± 0.4, estimated from the application of the maximum magnitude versus rate of decline (mmrd) relation [2] to the light curve of the 1966 outburst [3]. the presence of a clumpy ring with an inclination of 30-40 degrees, embedded within a spherical shell, was also recently reported using hubble space telescope observations during the nebular phase [1]. a variation of the orbital period after the 2011 outburst was detected, indicating an ejected mass in the outburst of 3 × 10−5 m� [4]. also, the orbital period was found to have increased during the long quiescence period between the last two outbursts of t pyx. moreover, with respect to the 1966 epoch, when no advanced x-ray and radio detectors were available, we have now a good knowledge of the emission of t pyx at these frequencies during its outburst [5]. in the previous work [6], we have presented the early spectral evolution, up to two months after the initial outburst. we have shown the hybrid spectral transition he/n fe ii he/n and initial ejecta velocities. in this work, we present the optical observations of the nebular phase, which was marked by the onset of the supersoft x-ray source (sss) emission which happened during the seasonal gap [7]. the optical observations presented here were obtained with the high-resolution spectrograph sarg, mounted on the 3.5m telescope nazionale galileo (tng) located at la palma. spectra with power resolution of r = 29000 were obtained but the complete sample, which includes also observations obtained at the very large telescope, will be presented in ederoclite et al. in preparation. then we briefly discuss the x-ray emission and finally we provide an updated scheme of the binary system configuration in t pyx, taking into account the recent results. 2 the nebular phase of t pyx the nebular phase in novae is easily recognized by the absence of absorption lines and the presence of forbidden emission lines, e.g. [oiii] and [nii], in the optical spectra, similarly to planetary nebula cases. indeed, these emission lines are due to photo-ionization of the ejecta by the central hot white dwarf (wd). at the same time, forbidden transitions originate mainly in rarified environments, so that any structure observed in corresponding line profiles is mainly due to the distribution of the material and not to self-absorption mechanisms. the analysis of the [oiii] 4959-5007 and [nii] 5755 λλ profiles can consequently provide hints on the geometry of the ejecta. these line profiles are well described by a large flat top base (vexp,1 ≈ 1700 km/s) overlapped by a saddleshaped and brighter emission, with vexp,2 ≈ 600 km/s, see fig. 1 for the case of the [oiii] 5007 λ. the high resolution allows us to detect a castellated structure on 264 http://dx.doi.org/10.14311/app.2015.02.0264 on the 2011 outburst of the recurrent nova t pyxidis the top of the profiles which is due to the presence of a clumpy ejecta structure. the saddle-shaped configuration suggests a non-spherical geometry of the ejecta, as generally discussed in [8] and further developed in several publications. in order to estimate the asphericity from forbidden lines we have used a method first presented in [9]. an analytic expression for the line profiles i(l) from an anisotropic distribution of the nova ejecta was provided: i(l) = k ( c v0 )α+1 ×[ lα−22 − l α−2 1 α − 2 l2(1 − 3cos2i) + lα2 − lα1 α (1 − sin2i) ] , l ≤ l1;[ lα−22 − l α−2 α − 2 l2(1 − 3cos2i) + lα2 − lα α (1 − sin2i) ] , l ≥ l1; (1) where c is the speed of light, v0 is the velocity at the inner radius of the expanding shell (we have assumed the maximum velocity observed for the bright profile, v0 ≈ 600 km/s), α describes the variation of the shell’s velocity with the distance form the wd, and l = (λ − λ0)/λ0 = v/c is our variable. the quantity -0.006-0.004-0.002 0.000 0.002 0.004 0.006 0 5.´10-13 1.´10-12 1.5´10-12 displacement l from @oiiidλλ 5007 fl ux he rg �c m 2 �s �a ng l figure 1: the saddle-shaped profile in the [oiii] 5007 line and the best fit (red line) obtained with the method described in the text. l2 = v2/c corresponds to the zero intensity along the wing of the line profile. a grid-search algorithm was developed to find the best fit for the inclination angle i that the binary orbital plane forms with the line of sight. for all the three lines, we find a large angle i ≈ 60 ± 10 degree, see also tab. 1, which contrasts some estimation of the i given before the 2011 outburst [14], and suggests a larger value for the binary inclination angle, see also [1]. we note that this approach does not take into account the presence of a clumpy structure. a more detailed approach is needed, and will be presented elsewhere. 3 the configuration of the binary system it is clear from the nebular spectra that the emission in t pyx is well described by quasi-spherical ejecta, which give rise to the larger flat top profile, and an equatorial ring of material, which is slower and gives rise to the saddle-shaped profiles. the origin of this equatorial ring is unknown: it is possible that it formed during the interaction of the main ejecta with pre-existing gas, which is equatorially distributed around the binary system (an hypothesis presented in [13]); another possibility can be that the early interaction of the ejecta with the secondary star could influence the subsequent expansion [10], but other mechanisms can also work. in all cases, it is necessary to shed light on the observed asphericity in t pyx, and the first ingredient we need is the mass of the wd. table 1: fit results of the line profiles of [oiii] 4959, 5007 and [nii] 5755, using the approach developed in [9]. ident. wavelength ( å) i (degree) v1 (km/s) α k (erg/cm 2/s/ang) 4959 59 ± 10 465 ± 6 -2.2 ± 0.3 (3.7 ± 0.1) × 10−15 5007 59 ± 10 477 ± 6 -2.3 ± 0.3 (1.09 ± 0.03) × 10−14 5755 60 ± 10 444 ± 6 -2.8 ± 0.3 (6.8 ± 0.5) × 10−16 265 l. izzo et al. � � � � � � � � figure 2: the location of the t pyx parameters (green stars) with respect to the m 31 nova correlations presented in [20]: (1) the ton-toff correlation, for t pyx measured from the time of maximum brightness; (2) the kt-toff relation, with kt given by the average value kt=55.7 ev; (3) the relation between ton and t2 and (4) the correlation between ton and the expansion velocity as measured from the hα emission line. x-ray observations of the sss phase can in principle provide this physical quantity. the sss emission is caused by stable hydrogen burning in the matter that was not ejected during the nova outburst. this burning matter forms an envelope around the wd and can be observed once the ejected material becomes optically thin to soft x-rays. the duration of the sss phase is inversely related to the wd mass [11], since more massive wds need to accrete less matter to trigger an outburst. given its recurrent nature, t pyx is expected to contain a massive wd. the sss phase of t pyx was first detected about 90 days after the outburst [7], when an increase in soft x-rays was observed by the swift-xrt [12]. t pyx remained in this state for ≈ 200 days, although around day 190 the luminosity started to decline. we do not enter here into a discussion on the details of this emission, which will be described in a forthcoming work. what is interesting in the context of this work is the relatively short duration of the sss phase, which suggest a large wd mass. observations obtained before 2011 reported a wd mass of 0.7 m� and a mass ratio of q = 0.2 [14], which implies a secondary star of 0.14 m�. the orbital period was observed with high accuracy to be porb = 1.8295 hrs. with some assumption on the type of the companion star, we can infer the mass of the secondary which overfills its roche lobe: ∆m2 = ρ2 4 3 π(∆r2) 3, where ρ2 is the density and ∆r2 the dimensions of the overfilling envelope of the secondary. from the q value and the absence of signatures of the secondary in optical spectra during the quiescent phase [3], we can assume that it is a low-mass m-type star. its density can be well approximated using the eggleton formula ρ2 = 107porb(hrs) −2 g cm−3 [15], while its radius can be derived from the patterson et al. formula r2 = 0.62m 0.61 2 r� [16]. the roche lobe size of the secondary is related to the q parameter through l2 = 1.631q 1/3m 1/3 1 porb(hrs) 2/3 1010 cm [17]. the result is that the envelope overfilling the secondary roche lobe is related to the wd mass. from the recent estimate of the mass ejecta after the 2011 266 on the 2011 outburst of the recurrent nova t pyxidis outburst (3 × 10−5m�, [4]), and assuming that every observed t pyx outbursts ejected similar masses, we obtain a total mass ejected from t pyx of about 2 × 10−4 m�. if this mass was accreted onto the wd from the secondary we can obtain a lower limit on the wd mass from the above formula. this results in mwd ≥ 0.8 m�. 4 final considerations our last result requires some important remarks. first, it was proposed that the mass ejected in t pyx is larger than the one accreted [3]. we do not consider this possibility in this work, i.e., that all the accreted mass results to be the ejected one. we know that this is not exactly true, since the sss phase is due to the burning of accreted hydrogen which is not ejected in the nova outburst. the mass of this burning envelope can be estimated assuming that the energy source for the sss phase is the gravitational contraction [18], but this contribution can be neglected. second, there is the possibility that a fraction of the matter overfilling the secondary’s roche lobe would flow out from the external lagrangian point l2, leading to the formation of an asymmetric structure around the binary system. such a mechanism would operate only if the primary roche lobe has been filled by accreted gas, leading to the detection of transient heavy elements in absorption (thea) during the early phases of the outburst [19], whose presence, in this case, is not certain. moreover, in tight systems where the secondary is less massive than the wd (q < 1) also centrifugal forces can contribute to such an outflow. there is again observational evidence that t pyx shares many properties common to other novae. as an example of this, we compare t pyx to the unbiased sample of m31 novae [20], in particular with respect to the correlations between some optical and x-ray observables: the blackbody temperature of the sss emitter, the turn-on and turn-off time of the sss emission, the expansion velocity of the ejecta and the t2 determined from the r-band light curve [21], see fig. 2. we see that t pyx lies in the same region occupied by classical novae observed in m31. all these results need to be figured in a unique picture for t pyx and can shed more light on cv systems with a mass ratio of less than unity. further results and conclusions will given in izzo et al. (2015) in preparation. acknowledgement we are very grateful to r. williams, s. n. shore, e. mason, d. de martino and r. gilmozzi for precious discussions. m.h. acknowledges support from an esa fellowship. references [1] sokoloski, j. l., crotts, a. p. s., uthas, h., et al.: 2013, apj 770, 33. [2] della valle, m. & livio, m.: 1998, apj 506, 818. [3] selvelli, p., cassatella, a., gilmozzi, r., et al.: 2008, a&a 492, 787. [4] patterson, j., oksanen, a., monard, b., et al.: 2013, arxiv/1303.0736. [5] nelson, t., chomiuk, l., roy, n., et al.: 2012, apj, 785, 78. [6] izzo, l., et al.: 2012, mem. s.a.it. s. 83, 830. [7] schwarz, g. j., osborne, j. p., page, k., et al.: 2011, atel 3754. [8] payne-gaposchkin, c. h.: 1957, the galactic novae. [9] vainu bappu, m. k. & menzel, d. h.: 1954, apj 119, 508. doi:10.1086/145857 [10] hutchings, j. b.: 1972, mnras 155, 357. doi:10.1093/mnras/155.3.357 [11] starrfield, s., truran, j. w., sparks, w. m., et al.: 1991, in extreme ultraviolet astronomy, ed. r. malina & s. bowyer, newyork, pergamon, p. 168. [12] gehrels, n., et al.: 2004, apj, 611, 1005. doi:10.1086/422091 [13] williams, r.: 2012, aj, 144, 98. [14] uthas, h., knigge, c. & steeghs, d.: 2010, mnras, 409,237. doi:10.1111/j.1365-2966.2010.17046.x [15] eggleton, p. p.: 1983, apj, 268, 368. doi:10.1086/160960 [16] patterson, j., kemp, j., harvey, d. a., et al.: 2005, pasp, 117, 1204. doi:10.1086/447771 [17] paczynski, b.: 1971, ara&a, 9, 183. doi:10.1146/annurev.aa.09.090171.001151 [18] krautter, j., oegelman, h., starrfield, s., et al.: 1996, apj, 456, 788. [19] williams, r., mason, e., della valle, m., et al.: 2008, apj, 685, 451. doi:10.1086/590056 [20] henze, m., pietsch, w., haberl, f., et al.: 2014, a&a, 563, a2. [21] della valle, m. & livio, m.: 1995, apj, 452, 704. 267 http://dx.doi.org/10.1086/145857 http://dx.doi.org/10.1093/mnras/155.3.357 http://dx.doi.org/10.1086/422091 http://dx.doi.org/10.1111/j.1365-2966.2010.17046.x http://dx.doi.org/10.1086/160960 http://dx.doi.org/10.1086/447771 http://dx.doi.org/10.1146/annurev.aa.09.090171.001151 http://dx.doi.org/10.1086/590056 l. izzo et al. discussion christian knigge: the system parameter we estimated in uthas et al. 2011 should be used with caution. they are based on a radial velocity study, and parameters resulting from such studies are known to be biased quite often. christian knigge: you consider r2 and l2 differing by ≈ 20%. are you really saying this could be the physical configuration, i.e. that the donor is overflowing the roche lobe by this amount ? luca izzo: the results that i’ve presented comes out naturally from formulae and physical considerations based on fixed binary parameters, as mwd = 0.7 m�, q = 0.2, porb = 1.8295 hrs, and considering the donor as a low mass m-type star. following a work by patterson et al., we obtain the r2 value, while from paczynski the value of l2. their interpretation depends obviously on a binary parameter as q presented in the uthas et al. work. anyway, other works consider that value of q as the actual one working in t pyx case. 268 introduction the nebular phase of t pyx the configuration of the binary system final considerations 139 acta polytechnica ctu proceedings 2(1): 139–142, 2015 139 doi: 10.14311/app.2015.02.0139 high-speed photo-polarimetry of magnetic cataclysmic variables s. b. potter1 1south african astronomical observatory, po box 9, observatory 7935, cape town, south africa corresponding author: sbp@saao.ac.za abstract i review recent highlights of the saao high-speed photo-polarimeter (hippo) on the study of magnetic cataclysmic variables. its high-speed capabilities are demonstrated with example observations made of the intermediate polar ny lup and the polar igrj14536-5522. keywords: cataclysmic variables polars intermediate polars optical polarimetry photometry x-rays individual: ny lup, igrj14536-5522. 1 introduction saaos hippo was designed and built in order to replace its highly successful but aging single channel equivalent, namely the uct (university of cape town) photo-polarimeter (cropper 1985). its purpose is to obtain simultaneous all-stokes parameters, of unresolved astronomical sources. in addition, it is capable of high speed, simultaneous 2 filtered, photo-polarimetry in order to permit investigations of rapidly varying polarized astronomical sources. of particular interest are magnetic cataclysmic variables (mcvs). in the following two sections we demonstrate its capabilities. 2 photo-polarimetry the intermediate polar ny lup figure 1: i-band photo-polarimetric observations of the intermediate polar ny lup. bottom and top panels are the phase-spin-folded photometry and circular polarization respectively. left and right panels were made with the 1.9m of the saao and the vlt respectively. figure reproduced from potter et al. (2012). 139 http://dx.doi.org/10.14311/app.2015.02.0139 s. b. potter fig. 1 shows two sets of i-band observations of the intermediate polar ny lup, phase-folded on its white dwarf spin period of 693s. the data in the left hand panels were made with the 1.9m telescope of the saao (taken from potter et al 2012) and the right hand panels were made with the 27 times larger vlt of the eso (taken from katajainen et al 2010). the saao polarimeter (hippo; potter et al 2010) is equiped with photo-multiplier tubes whereas the vlt polarimeter used a superior, 4 times more quantum efficient, ccd camera. both data sets show a spin modulation in the photometry and agree on the average value of circular polarization of ∼ -1.5 percent. however, contrary to expectations, the hippo observations are by far the superior dataset for the following reasons: the vlt photometry appears to have a residual repeating pattern of about 8 cycles per spin period, probably arising as a result of the different waveplate position angles. the hippo circular polarization shows a clear spin modulation whereas the vlt polarimetry displays a more random scatter about the mean. in addition the hippo also obtained simulatenously b-band observations. these observations (not shown) clearly show the presence of a circularly polarised spin modulation. the vlt b-band were not only not simultaneous, but also did not detect a spin modulation. furthermore the hippo simultaneously measured linear polarization (not shown) albiet a non-detection. the superior performance of the hippo is because it is optimised to measure polarization variability. specifically the exposure readout times of the photomultiplier tubes are very fast (sub-second) thus enabling polarization measurements to be made on a timescale much shorter than the intrinsic variability of the polarization. the ccd readout times on the vlt were too slow leading to the smearing of the polarimetric variability and systematic residuals in the photometry. 3 photo-polarimetry the polar igrj14536-5522 igrj14536-5522 (=swift j453.4-5524) was discovered as a hard x-ray source by integral (kuiper, keek, hermsen, jonker & steeghs 2006) and by swift/bat (mukai et al. 2006). a pointed swift/xrt observation led to the identification with a rosat all-sky survey (rass) source 1 rxs j145341.1-552146, and hence to its optical identification (masetti et al. 2006). based on these observations and the presence of short period optical and x-ray periodic modulations, revnivtsev et al. (2008) classified it as an ip. followup photo-polarimetric observations with the hippo (potter et al. 2010) however clearly shows orbitaly modulated photometry and circularly polarization from ∼ 0 to ∼ 18 per cent, unambiguously identifying igrj14536-5522 as a polar (fig.2, left and right upper panels respectively: from potter et al. 2010). in addition to the orbital modulation, a close inspection of the photometry and polarimetry reveals short period modulations throughout the orbit which, in addition to being a hard integral source, contributed to its initial mis-identification as an ip. the short period modulations are mostly consistent with being due to noise or flickering. however, detailed fourier analysis of all of the data reveals that some of the data sets show significant singularly persistent peaks that are consistent with qpos. the results are shown as a trailed amplitude spectrum in the second row of fig.2. between phases 0.2-1.0 the amplitude spectra do not show any significant peaks which indicates that the variations are mostly flickering or noise. however, there is a significant dominating signal centered on 0.0032(1) hz (5.2 minutes, indicated by the dashed line) between phases 1.0 and 1.3 in both the photometry and the polarimetry. this is clear evidence of a qpo. in the third row of fig. 2 we show the normalised photometry and polarimetry during the phase range that is dominated by the qpo. over plotted are the least squares fit of the qpo frequencies. as one can see, the qpo is very well described by the single dominant frequency as found in the trailed amplitude spectra. the plots in the bottom row of fig. 2 show the corresponding amplitude spectra for the qpo dominated phase range. our polarimetric results unambigously demonstrate that the qpo emissions (photometric and polarimetric) originate from the accretion shock. ultimately, these are caused by variations in the accretion flow. the most natural place to modulate the accretion flow would be at the threading region, perhaps caused by instabilities as a result of the interaction of the accretion flow with the magnetic field. we are now using mhd modelling to investigate this possibility. 140 high-speed photo-polarimetry of magnetic cataclysmic variables figure 2: left plots, a-d: the photometry, the corresponding trailed amplitude spectra, the normalised photometry for the phase range 1.0-1.3 and its corresponding amplitude spectra respectively. the solid curve is the least squares fit using the frequency derived from the trailed spectra (dashed line). right plots: as in the left plots but for the circular polarization. figure reproduced from potter et al. (2010). 4 conclusions we have shown that the saao high-speed photopolarimeter (hippo) is capable of high-speed, multifiltered, simultaneous all-stokes observations. it is therefore ideal for investigating rapidly varying astronomical sources such as magnetic cataclysmic variables. in particular white dwarf spin modulations and quasi-periodic-oscillations (qpos). acknowledgement i thank the organsers for giving me the opportunity to present some of my recent work and for a great conference. references [1] cropper, m. s.: 1985, mnras 212, 709. doi:10.1093/mnras/212.3.709 [2] kuiper, l., et al: 2006, atel 684. [3] katajainen, s., et al.: 2010, apj 724, 165. doi:10.1088/0004-637x/724/1/165 [4] masetti, n. et al: 2006, atel 783. [5] mukai, k. et al: 2006, atel 686. [6] potter, s.b., et al: 2010, mnras 402, 1161. doi:10.1111/j.1365-2966.2009.15944.x [7] potter, s.b., et al: 2012, mnras 420, 2596. doi:10.1111/j.1365-2966.2011.20232.x [8] revnivtsev, m., et al: 2008, a&a 489, 1121. discussion christian knigge: do you have any speculative ideas already for the origin of the ∼ 6 minute qpos? 141 http://dx.doi.org/10.1093/mnras/212.3.709 http://dx.doi.org/10.1088/0004-637x/724/1/165 http://dx.doi.org/10.1111/j.1365-2966.2009.15944.x http://dx.doi.org/10.1111/j.1365-2966.2011.20232.x s. b. potter stephen potter: since the qpo is seen in polarized light then it must be associated with the accretion shock region. so best guess is that the qpo may represent an instability in the threading region causing quasi-modulated accretion onto the white dwarf. pieter meintjes: in one of your slides you showed distinct qpo features with periods of ∼ few seconds. any idea what may have caused that and is that associated with high or low states in accretion? in other words, are they sometimes more prominent than other times (depending on accretion) and is the polarization level also fluctuating depending on accretion? stephen potter: photometric qpos of the order of a few seconds are thought to arise from oscillations in the shock itself. there is not yet a sufficient amount of observational data to see if there is any correlation with e.g. accretion rate and/or magnetic field strengths etc. we detect polarized qpos at ∼ minute timescales but not ∼ second timescales. 142 introduction photo-polarimetry the intermediate polar ny lup photo-polarimetry the polar igrj14536-5522 conclusions 222 acta polytechnica ctu proceedings 2(1): 222–225, 2015 222 doi: 10.14311/app.2015.02.0222 early super soft source spectra in rs oph j.-u. ness1 1xmm-newton science operations centre, esa, po box 78, 28691 villanueva de la cañada, madrid, spain corresponding author: juness@sciops.esa.int abstract recent swift x-ray monitoring campaigns of novae have revealed extreme levels of variability during the early super-softsource (sss) phase. the first time this was observed was during the 2006 outburst of the recurrent nova rs oph which was also extensively covered by grating observations with xmm-newton and chandra. i focus here on an xmm-newton observation taken on day 26.1, just before swift confirmed the start of the sss phase, and a chandra observation taken on day 39.7. the first observation probes the evolution of the shock emission produced by the collision of the nova ejecta with the stellar wind of the companion. the second observation contains bright sss emission longwards of 15 å while at short wavelengths, the shock component can be seen to have hardly changed. on top of the sss continuum, additional emission lines are clearly seen, and i show that they are much stronger than those seen on day 26.1, indicating line pumping caused by the sss emission. the lightcurve on day 39.7 is highly variable on short time scales while the long-term swift light curve was still variable. in 2007, we have shown that brightness variations are followed by hardness variations, lagging behind 1000 seconds. i show now that the hardness variations are owed to variations in the depth of the neutral hydrogen column density of order 25%, particularly affecting the oxygen k-shell ionization edge at 0.5 kev. keywords: cataclysmic variables recurrent novae spectroscopy x-rays individual: rs oph. 1 introduction the 2006 outburst of the recurrent symbiotic nova rs oph has attracted a large number of observers to study the outburst in unprecedented detail and many wavelength bands. in particular the coverage in x-rays has considerably improved since the previous outburst in 1985 by initiating the first high-density x-ray monitoring of a nova with the x-ray telescope (xrt) on board swift, starting on day 3.38 after initial explosion (bode et al. 2006a). early shock emission originated from kinetic energy from the nova ejecta that were dissipated in the slow, dense stellar wind of the giant companion. the expected x-ray spectrum is that of a collisional plasma which was confirmed by spectral models to the low-resolution exosat spectra taken in 1985 (figure 3 in (o'brien et al. 1992)) and the swift/xrt spectra (bode et al. 2006b). the x-ray grating spectrometers on board xmm-newton (reflection grating spectrometer rgs) and chandra (lowand high energy transmission grating spectrometers letgs and hetgs) allow spectral lines to be resolved, and in simultaneous rgs/hetgs spectra taken on day 13.8 after the initial explosion (2006 february 12.83), bremsstrahlung continuum, h-like and he-like emission lines, and numerous fe lines could be identified. a detailed analysis exploring the information from the emission lines yielded the distribution of electron temperatures and abundances (ness et al. 2009). further grating observations were taken, e.g., on days 26.1, 39.7, 54.0, 66.9, and 111.7 that were guided by continued dense swift monitoring (see table 1 in ness et al. 2009 and top of figure 2 in osborne et al. 2011). during the xmm-newton observation on day 26.1, a jump in count rate was discovered by (nelson et al. 2008) that was exclusively attributed to additional soft emission, possibly indicating the start of the sss phase. the bremsstrahlung continuum has faded and become softer, and also the ratios of h-like to he-like emission line strengths have shifted in favor of the he-like lines, indicating a cooling trend in the shocked plasma (ness et al. 2009). the swift observations of the ∼ 60 day sss phase are described in (osborne et al. 2011). between days ∼ 30 and 45, the x-ray count rate was highly variable between ∼ 10 counts per second (cps) and 200 cps, and then stabilized at ∼ 300 cps (figure 2 in osborne et al. 2011). this was a new discovery and was later also observed in other novae and might be a general phenomenon. during this high-amplitude phase, a chandra observation was taken on day 39.7, without yet knowing about the risks of observing during times of fainter emission. while this was the case, the soft emission was still bright enough for a well-exposed grating spectrum that was first presented by (ness et al. 2007). the blackbody-like continuum contained deep absorp222 http://dx.doi.org/10.14311/app.2015.02.0222 early super soft source spectra in rs oph tion lines from highly ionized species such as o viii and n vii that were blue shifted by ∼ 1200 km s−1. the line profiles contained clear signs of emission lines in the red wings (figure 5 of ness et al. 2007) which could either come from residual shock emission or are part of p cyg profiles. during this observation, the nova was also highly variably on shorter time scales, and (ness et al. 2007) reported that the hardness variations followed the same upand down trends but lagged 1000 seconds behind (see fig. 1). figure 1: reproduction of figure 8 from (ness et al. 2007). the chandra light curve on day 39.7, taken during the early high-amplitude variability phase, was also highly variable on shorter time scales. the hardness ratio, shown in the bottom, evolved with the same variability patterns but lagged 1000 seconds behind the brightness variations. reproduced by permission of the aas. i focus here on two new aspects: 1) the emission line components on top of the sss continuum (shown in figure 5 of ness et al. 2007), compared to the same lines before the sss phase started, are much stronger than expected from a cooling plasma (sect. 2.1). 2) closer inspection of the brightness/hardness changes during the day 39.7 observation show that the hardness changes are due to changes in the depth of the o i absorption edge. modeling shows that the overall column density increases with decreasing brightness (sect. 2.2). 2 observations and analysis a full log of all xmm-newton and chandra grating observation that were taken of the 2006 outburst of rs oph is given in table 1 in (ness et al. 2009) out of which i focus on the ones taken with xmm-newton between 2006 march 10, 23:04 and march 11, 02:21 (obsid 0410180201) and with chandra 2006 march 24, 12:25 15:38 (obsid 7296). the calibrated spectra are shown in direct comparison in fig. 2 using the same flux units without rescaling. figure 2: comparison of a pre-sss spectrum of rs oph taken on day 26.1 and an sss spectrum taken on day 39.7. in the top panel the entire wavelength range is shown in logarithmic units, illustrating that at short wavelengths, below ∼ 15 å, the spectra are similar, only showing a decline the bremsstrahlung continuum component from day 26.1 to day 39.5. in the panels below, narrow wavelength regions are shown where the pre-sss spectrum is added to the median of the day 39.7 spectrum in the same flux units. while short-wavelength lines have hardly changed, the excess emission lines on top of the sss continuum are much stronger than before the start of the sss phase. 2.1 contributions of shock emission to sss continuum spectrum in the top panel of fig 2, the full spectra are shown in logarithmic units to overcome the great contrast in brightness between the two observations. at wavelengths <∼ 15 å, the two spectra are similar in nature, mainly differing in the brightness of the continuum. the weaker continuum on day 39.7 can be explained by the continuation of the fading trend of the shocks that has already been established from the observations between days 13.8 and 26.1. the strengths of the emission lines have hardly changed. at wavelengths >∼ 15 å, the emission lines seen on day 26.1 were outshone by the bright continuum on day 39.7, and it is not intuitively clear how they have evolved. in the panels below, four lines are shown in more detail in linear units. the open histograms in the panels below represent the day 39.7 spectrum while the colored shades are the day 26.1 spectrum, added to the median flux of the day 39.7 spectrum, taken over the narrow wavelength range selected 223 j.-u. ness for each panel. the mg xii/xi lines have faded from day 26.1 and the ne x line at 12.1 å is about equally strong at both times while these lines have become somewhat narrower. the panels further below show that the lines on top of the sss continuum have become stronger from day 26.1 to 39.7, depending on the strength of the continuum. the fe xvii line at 15 å, in the wien tail of the sss continuum, is only slightly stronger while the o viii and n vii lines, near the peak of the sss continuum, are much stronger, defying the general trend of fading emission from the shocks. a similar result has been found by (schönrich & ness 2008) who show in their figure 1 the evolution of the volume emission measure for various emission lines as a function of their peak formation temperature assuming collisional equilibrium. the line fluxes of o vii and o viii from the observation taken on day 39.7 yield clearly discrepant values, orders of magnitude brighter than in observations without sss emission, while those emission lines that arise at shorter wavelengths are consistent with the other observations. 2.2 spectral changes with variability during early sss phase the high-amplitude variations during the early sss phase are still not understood, leaving all options open such as changes in absorption, in intrinsic brightness, temperature, or in the rate of mass loss. a first approach to narrow down the options, i present the spectra and how they changed with variability. during the early variability phase of rs oph, a short chandra observation taken (day 39.7), and the spectral evolution is illustrated in fig. 3. the light curve, already shown in fig. 1, is shown in the right, turned around by 90o clockwise to follow the vertical time axis in downward direction. the blue dotted line is the hardness light curve from the bottom panel of fig. 1 that was rescaled to fit in the same graph. along the same vertical time axis, a series of adjacent letgs spectra are shown in the central panel with wavelength along the horizontal axis and flux encoded in a color scheme, where increasing flux is represented by colors light green, yellow, orange, red, dark blue to light blue. in the top panel, two spectra are shown that have been integrated over the time intervals marked with horizontal dashed lines in the central panel shaded areas in the right panel. the colors of the dashed lines and shades correspond to the plot style of the spectra in the top, thus light blue and orange shadings in the right correspond to the light blue shade and orange thick line in the top, respectively. the two spectra have been normalized to coincide in the wavelength range 15 − 22 å. ness et al. (2007) had extracted spectra from time intervals of bright and faint episodes, roughly corresponding to the time intervals 0.3 − 1.0 hours of elapsed time (bright) and 1.1 − 2 hours, respectively, and found differences in the spectral shape. i have now chosen time intervals that are shifted by 1000 sec in order to probe the hardness light curve rather than the intensity light curve. the two normalized spectra in the top clearly show that the variability in hardness is purely owed to changes in the depth of the o i edge at 22.8 å, yielding the steeper edge 1000 sec after a lowflux episode. ness et al. (2007) had found changes in o i on longer time scales between the grating observations taken on days 39.7, 54.1, and 66.9 and explained these by changes in the degree of ionization of oxygen. only neutral oxygen produces the deep edge at 22.8 å while ionized oxygen is more transparent to x-rays, leading to a flatter edge. the intense soft continuum x-ray source can ionize the neutral elements in the surroundings and thus make them more transparent. on the other hand, if the continuum source fades, the surrounding material recombines, leading to an increase in the o i edge. figure 3: dynamic spectrum for the chandra observation of rs oph taken on day 39.7 after the 2006 outburst. the zero-th order light curve is shown to the right with three colored shades marking time intervals from which the spectra in the top have been extracted. line labels that belong to strong transitions are included. the central image is a brightness spectral/time map that uses a color code from light green to light blue representing increasing flux values. the changes in the depth of the o i absorption edge could be caused by changes in the degree of ionization of circumstellar oxygen. ness et al. (2007) argued that the observed time lag of 1000 seconds is consistent with the time scale for ionization/recombination in a plasma with density ∼ 1011 cm−3. since all circumstellar material would experience the same changes in their degree of ionization, this would lead to an overall change in nh. another possibility is a non-uniform radial distribution of the oxygen abundance within the ejecta that could be caused by beta decay of oxygen isotopes produced during cno burning. if during times of brigher and fainter emission, different regions of the photosphere are visi224 early super soft source spectra in rs oph ble, the difference in the oxygen abundance would only affect the depth of the o i absorption edge. in fig. 4, i test these two scenarios in the top and bottom panels, respectively. shown is the normalized faint/soft spectrum in orange shadings (orange line in the top panel of fig. 3) while the black curve represents a modification of the brighter/harder spectrum, trying to reproduce the faint/soft spectrum. in the top panel of fig. 4, the black curve was computed by dividing the bright/hard spectrum itself by the transmission coefficients calculated from the warm absorption model developed by [?], and then normalized. the faint/soft spectrum is qualitatively well reproduced assuming a value of nh = 6 × 1020 cm−2, indicating that the change in hardness can be explained by an increase of nh by this amount. the full amount of the hydrogen column was found by [?] as nh = nh,ism + nh,cs = 4.7 × 1021 cm−2, consisting of the well-known interstellar component nh,ism = 2.4 × 1021 cm−2 and an additional circumstellar component nh,cs. an increase by 6 × 1020 cm−2 thus corresponds to ∼ 26% of nh,cs. figure 4: testing two hypotheses to explain the spectral changes between hard/bright (blue shades in top panel of fig. 3) and soft/faint (orange thick line in fig. 3) spectra by the black thin lines. the normalized soft/faint spectrum is shown as orange shadings while the black line corresponds to two types of modifications of the bright/faint spectrum, also normalized (see text). top: overall increase of nh by 26% of nh,cs. bottom: increase only of o i column by changes in o abundance of 15%. in the bottom panel of fig. 4, the black curve was computed by first dividing the bright-hard spectrum by the transmission coefficients of an absorber with nh = 4.7 × 1021 cm−2, assuming cosmic abundances and then multiplying by the coefficients resulting from the same absorption model assuming a modified oxygen abundance. the faint/soft spectrum is well reproduced up to ∼ 27 å, but the black curve drops well below the faint/soft spectrum at longer wavelengths. the better agreement between the black line and the faint/soft spectrum in the upper panel demonstrates that an overall increase of nh has occurred rather than a change in the oxygen abundance alone. 3 summary and conclusions a comparison of grating x-ray spectra before and after the start of the sss phase shows that the emission lines that were seen on top of the sss continuum are not simply a continuation of the early shock emission. only emission lines that arise at wavelengths were strong continuum emission is present are amplified compared to the pre-sss level, strongly suggesting photoexcitation effects. the line profiles in the early sss spectrum of rs oph can thus be understood as p cyg profiles. the early high-amplitude variability in rs oph and other novae still awaits an explanation. the chandra observation during a minimum of these variations revealed that variability also occurs on shorter time scales. the comparisons in this article show that the hardness changes following brightness changes are consistent with variations in nh of ∼ 26% which can be explained by variations in the overall degree of ionization caused by the changes in intensity and thus variations in the effectiveness of photoionization. references [1] bode, m. f., obrien, t. j., davis, r. j., et al. 2006a, iaucirc, 8675, 2 [2] bode, m. f., obrien, t. j., osborne, j. p., et al. 2006b, apj, 652, 629 doi:10.1086/507980 [3] nelson, t., orio, m., cassinelli, j. p., et al. 2008, apj, 673, 1067 doi:10.1086/524054 [4] ness, j.-u., starrfield, s., beardmore, a., et al. 2007, apj, 665, 1334 doi:10.1086/519676 [5] ness, j.-u., drake, j. j., starrfield, s., et al. 2009, aj, 137, 3414 [6] o'brien, t. j., bode, m. f., & kahn, f. d. 1992, mnras, 255, 683 doi:10.1093/mnras/255.4.683 [7] osborne, j. p., page, k. l., beardmore, a. p., et al. 2011, apj, 727, 124 doi:10.1088/0004-637x/727/2/124 [8] schönrich, r. a. & ness, j.-u. 2008, in asp conf. series, 401, 291; rs ophiuchi (2006) and the recurrent nova phenomenon, ed. a. evans, m. f. bode, t. j. obrien, & m. j. darnley [9] wilms, j., allen, a., & mccray, r. 2000, apj, 542, 914 225 http://dx.doi.org/10.1086/507980 http://dx.doi.org/10.1086/524054 http://dx.doi.org/10.1086/519676 http://dx.doi.org/10.1093/mnras/255.4.683 http://dx.doi.org/10.1088/0004-637x/727/2/124 introduction observations and analysis contributions of shock emission to sss continuum spectrum spectral changes with variability during early sss phase summary and conclusions 286 acta polytechnica ctu proceedings 2(1): 286–290, 2015 286 doi: 10.14311/app.2015.02.0286 rotation of the mass donors in high-mass x-ray binaries and symbiotic stars k. stoyanov1, r. zamanov1 1institute of astronomy and national astronomical observatory, bulgarian academy of sciences, 72 tsarigradsko shousse blvd., 1784 sofia, bulgaria corresponding author: kstoyanov@astro.bas.bg abstract our aim is to investigate the tidal interaction in high-mass x-ray binaries and symbiotic stars in order to determine in which objects the rotation of the mass donors is synchronized or pseudosynchronized with the orbital motion of the compact companion. we find that the be/x-ray binaries are not synchronized and the orbital periods of the systems are greater than the rotational periods of the mass donors. the giant and supergiant high-mass x-ray binaries and symbiotic stars are close to synchronization. we compare the rotation of mass donors in symbiotics with the projected rotational velocities of field giants and find that the m giants in s-type symbiotics rotate on average 1.5 times faster than the field m giants. we find that the projected rotational velocity of the red giant in symbiotic star mwc 560 is v sin i= 8.2 ± 1.5 km s−1, and estimate its rotational period to be prot = 144 306 days. using the theoretical predictions of tidal interaction and pseudosynchronization, we estimate the orbital eccentricity e = 0.68 − 0.82. keywords: stars: binaries: symbiotic stars: rotation stars: late type stars: binaries: close x-rays: binaries. 1 introduction a high-mass x-ray binary system consists of a compact object (a neutron star or a black hole) accreting material from an o or b companion star. they are divided into be/x-ray binaries (main-sequence star as a companion) and giant/supergiant x-ray binaries (giant or supergiant star as a companion). accretion of matter is different for both types of x-ray binaries. in the be/x-ray binaries, the compact object crosses the circumstellar disc and accretes matter from that disk. in the giant/supergiant x-ray binaries, the mass donor ejects a slow and dense wind radially outflowing from the equator and the compact object directly accretes the stellar wind through bondi-hoyle-lyttleton accretion. symbiotic stars are interacting binaries, consisting of an evolved giant (either a normal red giant in s-types symbiotics or a mira-type variable in d-types symbiotics) transferring mass to a hot and luminous white dwarf or neutron star. the symbiotic stars are surrounded by a rich and luminous nebula resulting from the presence of both an evolved giant with a heavy mass-loss and of a hot companion abundant in ionizing photons and often emanating its own wind. 2 synchronization and pseudosynchronization in a binary with a circular orbit the rotational period of the primary, prot, reaches an equilibrium value at the orbital period, porb = prot. in other words the synchronous rotation (synchronization) means that the rotational period is equal to the orbital period. in a binary with an eccentric orbit, the corresponding equilibrium is reached at a value of prot which is less than porb, the amount less being a function solely of the orbital eccentricity e. in practice, in a binary with an eccentric orbit the tidal force acts to synchronize the rotation of the mass donor with the motion of the compact object at the periastron pseudosynchronous rotation. to calculate the period of pseudosynchronization, pps, we use (hut 1981): pps = (1 + 3e2 + 3 8 e4)(1 − e2) 3 2 1 + 15 2 e2 + 45 8 e4 + 5 16 e6 porb. (1) 2.1 stars with radiative envelopes following hurley, tout & pols (2002) the circularization timescale for stars with radiative envelopes is: 1 τcirc = 21 2 ( gm1 r31 )1 2 q2 (1 + q2) 11 6 e2 ( r1 a )21 2 , (2) 286 http://dx.doi.org/10.14311/app.2015.02.0286 rotation of the mass donors in high-mass x-ray binaries and symbiotic stars where m1 and r1 are the mass and the radius of the primary respectively, q2 is the mass ratio m2/m1, and a is the semi-major axis. the second-order tidal coefficient e2 = 1.592 ×10−9m2.841 . the synchronization time scale is given as, τsync = kτcirc, (3) where k is: k ≈ 0.015 rg 1 + q2 q2 ( r1 a )2 . (4) for the gyration radius of the primary rg we adopt rg ≈ 0.16 for giants, and rg ≈ 0.25 for main sequence stars (claret & gimenez, 1989). 2.2 stars with convective envelopes following hurley, tout & pols (2002) the synchronization timescale for stars with convective envelopes is: τsyn ≈ 800 ( m1r1 l1 )1/3 m21 ( m1m2 + 1)2 r61 p4orb yr, (5) where l1 is the luminosity of the giant. the circularization time scale is: 1 τcirc = 21 2 ( k t ) q2(1 + q2) ( r1 a )8 . (6) in eq. 6, (k/t) is:( k t ) = 2 21 fconv τconv menv m1 yr−1, (7) where renv is the depth of the convective envelope, menv is the envelope’s mass, and τconv = 0.4311 ( menvrenv(r1 − 12renv) 3l1 )1 3 yr (8) is the eddy turnover time scale (the time scale on which the largest convective cells turnover). the numerical factor fconv is fconv = min [ 1, ( ptid 2τconv )2] , (9) where ptid is the tidal pumping time scale given by 1 ptid = ∣∣∣∣ 1porb − 1prot ∣∣∣∣ . (10) the pseudosynchronization timescale is τps = (7/3(α 3)) τcirc, where α is a dimensionless quantity, representing the ratio of the orbital and rotational angular momentum: α = q2 1 + q2 1 r2g ( a r1 )2 . (11) for a red giant we adopt rg ≈ 0.3 (claret, 2007). in all equations, the masses, the radii and the lumunosities are in solar units. 3 high-mass x-ray binaries the orbital and stellar parameters of 13 high-mass xray binaries are given in table 1 and table 2 in stoyanov & zamanov (2009). we add 2 more objects 4u 2206+54 and mwc 148. the orbital and stellar parameters are taken from ribó et. al. (2006) and casares et al. (2012) respectively. using eq.2 and eq.3 we estimate the circularization and synchronization timescales. the results are given in table 1. the lifetime of a star on the main sequence can be estimated as τms = 10 10(m�/m) 2.5 years (hansen & kawaler, 1994). comparing these lifetimes with τsync from table 1, we see that among the be/x-ray binaries only for lsi+610303 is τsync ∼ τms. this is the only be/x-ray binary for which we can expect considerable changes of the rotation of the primary during the lifetime of the be star. the lifetime of the giant is comparable or longer then τcirc and τsync for the giant/supergiant systems with short orbital periods. the exceptions are v725 tau and bp cru, for which τsync and τcirc are longer than the lifetime of the giant/supergiant stage. on fig.1 in stoyanov & zamanov (2009) is plotted prot versus pps. the giant/supergiant systems are located close to the line pps = prot, while those with mass donors from spectral class v are far away from the equilibrium. in the be/x-ray systems bq cam, v635 cas, v725 tau and 4u 2206+54, the tidal force spinning down the donor star. for the system lsi+610303, the rotation of the mass donor is close to pseudosynchronization. this is the only be/x-ray binary in which τsync is comparable with the life-time of the binary. in the binaries x per and mwc 148, the neutron star is far away from the be star and the tidal force is weak. giant and supergiant systems are close to (pseudo)synchronization. in these binaries the rotation of the mass donors is influenced by the presence of the compact object. in lmc x-4 and cen x-3, the mass donors are synchronized and the orbits are circularized. with respect to the rotation of the mass donor, v725 tau is similar to the be/x-ray binaries. cyg x-1 is synchronized and almost circularized. v830 cen is pseudosynchronized but not circularized yet. the systems lsi+650010 and vela x-1 are close to pseudosynchronization and the tidal force accelerates the rotation of the mass donors. in the case of smc x-1, the tidal force acts as a decelerator of the rotation of the mass donor. in bp cru, a gas stream from the mass donor 287 k. stoyanov, r. zamanov exists, probably resulting from the strong tidal force and spin-up of the mass donor (leahy & kostka, 2008). table 1: calculated time scales. given here are the name of the object, synchronization time scale, circularization time scale, and the lifetime. object τsync [yr] τcirc [yr] lifetime [yr] be/x-ray binaries lsi+610303* 2.8×107 2.4×108 5.6×107 x per 6.2×1017 1.8×1021 1.1×107 bq cam 3.5×1011 7.6×1013 3.9×106 v635 cas 1.4×1011 9.5×1012 7.3×106 4u 2206+54 4.9×109 3.7×1011 7.3×107 mwc 148 1.2×1017 5.1×1020 9.8×107 giant systems v725 tau 2.8×1012 8.0×1014 4×105 lmc x-4 4.5×102 7.7×102 1×106 cen x-3 2.3×103 4.2×103 5×105 supergiant systems v830 cen 7.5×103 1.4×104 1×106 lsi+650010 1.3×104 3.9×104 1×106 vela x-1 1.0×104 2.8×104 3.9×105 smc x-1 3.3×104 8.2×104 8.8×105 bp cru 1.8×106 8.8×106 8×104 cyg x-1 < 1 < 1 1×105 *assuming neutron star as a secondary component. 4 s-type symbiotic stars 43 symbiotic stars have been observed with feros spectrograph at the 2.2m eso telescope of the la silla observatory (zamanov et al. 2007). the data for the rotation of 55 field red giants are taken from the literature. m giants in s-type symbiotics rotate faster than the field m giants. histograms of the available v sin i data for the red giants are plotted in fig.2 in zamanov & stoyanov (2012). for the field m0iii-m6iii giants we calculate a mean v sin i=5.0 km s−1, median v sin i=4.3 km s−1, and standard deviation of the mean σ = 4.0 km s−1. for the m0iii-m6iii giants in symbiotics, we get a mean v sin i= 7.8 km s−1, median v sin i=8.0 km s−1, and standard deviation of the mean σ =2.1 km s−1. there are 5 objects in our sample that deviate from the synchronization. these objects are rs oph, mwc 560, ch cyg, cd-43◦14304 and z and. in three of them collimated jets are detected: z and (skopal et al. 2009); ch cyg (crocker et al. 2002), mwc 560 (tomov et al. 1990). additionally to the jets, ejection of blobs are detected from rs oph and ch cyg (iijima et al. 1994). this confirms the suggestions that in the jet-ejecting symbiotics the mass donors rotate faster than the orbital periods. probably there is a link between the jets and the mass donor rotation. on fig.1 are plotted together the high-mass x-ray binaries and the s-type symbiotic stars. it shows that none of the objects in our sample is above the line of synchronization. figure 1: prot versus porb on a logarithmic scale. the blue circles indicates the high-mass x-ray binaries filled symbols for giant/supergiant systems and open symbols for be/x-ray binaries. the red triangles indicates the s-type symbiotic stars. 5 orbital eccentricity of mwc 560 mwc 560 is a symbiotic star, which consists of a red giant and a white dwarf (tomov et al. 1990). the most spectacular features of this object are the collimated ejections of matter with velocities of up to ∼ 6000 km s−1 (tomov et al. 1992) and the resemblance of its emission line spectrum to that of the low-redshift quasars (zamanov & marziani, 2002). the jet ejections are along the line of sight and the system is seen almost pole-on (i < 16◦). this makes it difficult to obtain the orbital eccentricity of the system in a conventional way. for the system we adopt rg = 140 ± 7 r�, lg ∼ 2400 l�, mg = 1.7 m�, mwd = 0.65 m�, porb = 1931 ± 162 day (gromadzki et al. 2007), renv = 0.9 rg 288 rotation of the mass donors in high-mass x-ray binaries and symbiotic stars and menv = 1.0 m� (herwig 2005). with the above values of the parameters assumed, we derive the semimajor axis of the orbit to be a ≈ 860 r�. using these parameters, we calculate from eq. 5 and eq. 6 the synchronization and circularization time scales: τsync = 2.6 × 104 yr and τcirc = 3.1 × 106 yr. the typical lifetime of a symbiotic star is τss ∼ 105 yr (yungelson et al., 1995). from the rate of accretion on the white dwarf, ṁacc ≈ 5 × 10−7 m� (schmid et al. 2001), we can estimate, that it will take 106 yr to accrete ∼ 0.5 m� from the envelope of the red giant companion. because the giant also losses mass via stellar wind, we find that the lifetime of the symbiotic phase of mwc 560 should be τss ≤ 106 yr. for mwc 560 we have therefore the situation in which τps < τsyn < τss < τcirc. this means that the symbiotic phase is long enough that the tidal forces can (pseudo)synchronize the rotation of the red giant. on the other hand, the value of τcirc demonstrates that the symbiotic lifetime of mwc 560 is shorter than the circularization time, and therefore the orbit can be eccentric. this is in agreement with the observational evidences found by fekel et al. (2007) that the symbiotic stars with porb > 800 days tend to have eccentric orbits. the above implies that in mwc 560, the red giant is probably synchronized, but the orbit is not circularized. to determine the orbital eccentricity of mwc 560, we need to calculate prot for the mass donor. we analyzed 21 high resolution spectra of mwc 560 and obtained value for v sin i= 8.2 ± 1.5 km s−1. using rg = 140 ± 7 r� and i = 12◦ − 16◦, we calculate prot = 144 − 306 days. this value is less than the orbital period. mwc 560 should be close to synchronization or pseudosynchronization, and prot = pps. using eq. 1 we can estimate the orbital eccentricity to be e = 0.68 − 0.82. 6 conclusions using rotational velocity measurements and the theory of synchronization/pseudosynchronization we: (1) find that the be/x-ray binaries are far away from (pseudo)synchronization. the tidal force in the be/x-ray binaries acts as a decelerator of the rotation of the mass donors. the only be/x-ray binary which is close to pseudosynchronization is the lsi+610303. the objects containing mass donors of spectral class i and iii typically have prot ∼ pps and are close to (pseudo)synchronization; (2) demonstrate that the m giants in symbiotic stars rotate faster than the field giants. most symbiotics with orbital period less than 1000 d are synchronized; (3) show that the high-mass x-ray binaries and the s-type symbiotic stars are either on the line of synchronization or they are under the line. none of the objects in our sample is above the line of synchronization. (4) calculate that the orbit of the symbiotic star mwc 560 should be highly eccentric, with e ∼ 0.7. acknowledgement we thank the anonymous referee for constructive comments. this work was supported by the op “hrd“, esf and bulgarian ministry of education, youth and science under the contract bg051po001-3.3.06-0047. references [1] casares, j., ribó, m., ribas, i., paredes, j. m., vilardell, f., & negueruela, i.: 2012, mnras, 421, 1103 doi:10.1111/j.1365-2966.2011.20368.x [2] claret, a.: 2007, a&a 467, 1389 [3] claret, a., & gimenez, a.: 1989, a&as, 81, 37 [4] crocker, m. m., davis, r. j., spencer, r. e., et al.: 2002, mnras, 335, 1100 doi:10.1046/j.1365-8711.2002.05705.x [5] fekel, f. c., hinkle, k. h., joyce, r. r., wood, p. r., lebzelter, t.: 2007, aj 133, 17 [6] gromadzki, m., miko lajewska, j., whitelock, p. a., marang, f.: 2007, a&a 463, 703 [7] hansen, c. j., & kawaler, s. d.: 1994, stellar interiors. physical principles, structure, and evolution, springer-verlag [8] herwig, f.: 2005, ara&a 43, 435 doi:10.1146/annurev.astro.43.072103.150600 [9] hurley, j. r., tout, c. a., & pols, o. r.: 2002, mnras, 329, 897 doi:10.1046/j.1365-8711.2002.05038.x [10] hut, p.: 1981, a&a, 99, 126 [11] iijima, t., strafella, f., sabbadin, f., & bianchini, a.: 1994, a&a, 283, 919 [12] leahy, d. a., & kostka, m.: 2008, mnras, 384, 747 doi:10.1111/j.1365-2966.2007.12754.x [13] ribó, m., negueruela, i., blay, p., torrejón, j. m., & reig, p.: 2006, a&a, 449, 687 [14] schmid, h. m., kaufer, a., camenzind, m., rivinius, t., stahl, o., szeifert, t., tubbesing, s., wolf, b.: 2001, a&a 377, 206 [15] skopal, a., pribulla, t., budaj, j., et al.: 2009, apj, 690, 1222 doi:10.1088/0004-637x/690/2/1222 [16] stoyanov, k. a., & zamanov, r. k.: 2009, astronomische nachrichten, 330, 727 289 http://dx.doi.org/10.1111/j.1365-2966.2011.20368.x http://dx.doi.org/10.1046/j.1365-8711.2002.05705.x http://dx.doi.org/10.1146/annurev.astro.43.072103.150600 http://dx.doi.org/10.1046/j.1365-8711.2002.05038.x http://dx.doi.org/10.1111/j.1365-2966.2007.12754.x http://dx.doi.org/10.1088/0004-637x/690/2/1222 k. stoyanov, r. zamanov [17] tomov, t., zamanov, r., kolev, d., georgiev, l., antov, a., mikolajewski, m., esipov, v.: 1992, mnras 258, 23 doi:10.1093/mnras/258.1.23 [18] tomov, t., kolev, d., zamanov, r., georgiev, l., antov, a.: 1990, nature 346, 637 doi:10.1038/346637a0 [19] yungelson, l., livio, m., tutukov, a., & kenyon, s. j.: 1995, apj 447, 656 doi:10.1086/175908 [20] zamanov, r. k., bode, m. f., melo, c. h. f., et al.: 2007, mnras, 380, 1053 doi:10.1111/j.1365-2966.2007.12150.x [21] zamanov, r., marziani, p.: 2002, apjl 571, l77 doi:10.1086/341367 [22] zamanov, r. k., & stoyanov, k. a.: 2012, bulgarian astronomical journal, 18, 41 290 http://dx.doi.org/10.1093/mnras/258.1.23 http://dx.doi.org/10.1038/346637a0 http://dx.doi.org/10.1086/175908 http://dx.doi.org/10.1111/j.1365-2966.2007.12150.x http://dx.doi.org/10.1086/341367 introduction synchronization and pseudosynchronization stars with radiative envelopes stars with convective envelopes high-mass x-ray binaries s-type symbiotic stars orbital eccentricity of mwc 560 conclusions 210 acta polytechnica ctu proceedings 1(1): 210–214, 2014 210 doi: 10.14311/app.2014.01.0210 electron acceleration in supernovae and millimeter perspectives keiichi maeda1,2 1department of astronomy, kyoto university, japan 2kavli institute for the physics and mathematics of the universe (wpi), university of tokyo, japan corresponding author: keiichi.maeda@kusastro.kyoto-u.ac.jp abstract supernovae launch a strong shock wave by the interaction of the expanding ejecta and surrounding circumstellar matter (csm). at the shock, electrons are accelerated to relativistic speed, creating observed synchrotron emissions in radio wavelengths. in this paper, i suggest that sne (i.e., ∼< 1 year since the explosion) provide a unique site to study the electron acceleration mechanism. i argue that the efficiency of the acceleration at the young sn shock is much lower than conventionally assumed, and that the electrons emitting in the cm wavelengths are not fully in the diffusive shock acceleration (dsa) regime. thus radio emissions from young sne record information on the yet-unresolved ‘injection’ mechanism. i also present perspectives of millimeter (mm) observations of sne – this will provide opportunities to uniquely determine the shock physics and the acceleration efficiency, to test the non-linear dsa mechanism and provide a characteristic electron energy scale with which the dsa start dominating the electron acceleration. keywords: acceleration of particles radiation mechanism: non-thermal supernovae: general. 1 introduction the most promising particle acceleration mechanisms require a strong shock wave, e.g., by the diffusive shock acceleration (dsa) mechanism where the particles acquire energy through repeated collisions between upand down-streams of a shock wave (fermi, 1949; blandford & ostriker, 1978; bell, 1978). supernova remnants (snrs) are believed to be the origin of cosmic rays at least up to ∼ 1015ev (e.g., bamba et al., 2003). there is one key issue in this picture for electrons – how the electrons are ‘pre-accelerated’. for the dsa mechanism to work efficiently, a particle must already have an enough kinetic energy. supernovae (sne), at the age of ∼< 1 year, also produce emissions which are believed to be originated by relativistic electrons, accelerated at a strong shock created by the expanding sn ejecta running into circumstellar matter (csm). radio emissions from sne are interpreted as the synchrotron emission, and x-rays from some sne have been suggested to be emitted through the inverse compton (ic) mechanism (e.g., chevalier & fransson, 2006 for a review). however, most of analyses on the non-thermal emissions from sne have been focusing on deriving the csm environment, rather than the acceleration mechanism (e.g., soderberg et al., 2012). in this paper, i argue that young sne provide a unique site to study the electron acceleration mechanism. i also suggest that millimeter (mm) observations, which are becoming feasible with new observatories like alma, can potentially provide essential information on this issue. 2 non-thermal emissions a situation around young sne related to the nonthermal emission is similar to that for snrs. energy transfer from the shock wave kinetic energy to relativistic particles and that to magnetic field are roughly described by equipartition (fransson et al., 1996). i adopt conventional notation – �e and �b describe constant fractions of the shock wave energy transferred to the relativistic electrons and the magnetic field, respectively. our arguments are based on modeling emissions from so-called striped envelope sne (se-sne; or sne iib/ib/ic) that are believed to be explosions of he or co stars (having lost at least the h envelope). in these sne, the radio emission is well described by synchrotron emissions with the synchrotron self-absorption (ssa) at low frequencies (see, e.g., chevalier & fransson, 2006). under some standard assumptions (björnsson & fransson, 2004; chevalier & fransson, 2006; maeda et al. 2012, 2013a), the synchrotron properties can be described by the following parameters: • p: power law index of spectral energy distribution of injected relativistic electrons. • m: power law index of shock evolution in time (r ∝ tm). 210 http://dx.doi.org/10.14311/app.2014.01.0210 electron acceleration in supernovae and millimeter perspectives table 1: characteristics of the synchrotron emission from young sne (lν ∝ ναtβ), for the adiabatic limit and for the synchrotron and ic cooling limits, respectively (maeda, 2013a). indices adiabatic syn. ic α 1−p 2 −p 2 −p 2 β (3m− 3) + 1−p 2 (3m− 3) + 2−p 2 (5m− 5) + 2−p 2 + δ α(p = 2) −1 2 −1 −1 β(p = 2) (3m− 3) − 1 2 (3m− 3) (5m− 5) + δ α(p = 3) −1 −3 2 −3 2 β(p = 3) (3m− 3) − 1 (3m− 3) − 1 2 (5m− 5) − 1 2 + δ • δ: power law index of optical/nir sn emission in time (l ∝ tδ). • a∗: csm density scale (ρcsm ∝ a∗r−2; normalized as a∗ ∼ 1 for ṁ ∼ 10−5m�yr−1 with the mass loss wind velocity of 1, 000 km s−1). • �e: efficiency of the electron acceleration. • �b: efficiency of the magnetic field generation/amplification. note that the shock evolution (m) is mainly determined by the csm density distribution, e.g., by a self-similar solution (chevalier, 1982). table 1 shows expected synchrotron properties, lν ∝ ναtβ. from the observed properties (α, β) one can almost uniquely determine the power law indices (p, m, δ). there is a degeneracy in the other parameters (i.e., in the ‘scales’). properties of the ssa-synchrotron are described by two characteristic observables (peak date and luminosity), while these are described by the three model parameters (i.e., a∗, �e, �b). 3 efficiency of electron acceleration figure 1 shows how one can constrain the shock microphysics and csm density. an example is given for intensively observed nearby sn iib 2011dh. thanks to detailed models of the optical emission (bersten et al., 2012), the sn ejecta properties (mass and energy) have been strongly constrained – model a adopts the shock wave dynamics expected from the optical emission model. model b is shown for illustration, which assumes the dynamics so that �e ∼ 0.1, but this fails to explain the optical behavior. adopting model a, �e cannot be as large as ∼ 0.1 which has been conventionally assumed, since such a situation requires extremely large mass loss rate (a∗). then the expected thermal emission in x-rays would be much stronger than observed. indeed, from the x-ray strength, a∗ ∼< 30, thus �e ∼< 0.01 must apply. also, from the energy conservation, a∗ ∼< 2 is rejected (otherwise �b > 0.3). from these arguments, 0.005 ∼< �e ∼< 0.01 and �b ∼> 0.001 are obtained as robust constraints. this also indicates that α ≡ �e/�b < 10. 0.1 1 10 1e-3 0.01 0.1 model b model a b e e figure 1: �e and �b derived for sn 2011dh, as a function of a∗ (maeda, 2012). so, a strong constraint can be placed on �e. there is another independent argument against a large value of �e. figure 2 shows the models with small �e and large �e. large �e should produce a detectable cooling effect in radio properties, which was however not detected. this argument on the ic cooling effect should apply to sne in general. i note that sometimes a large value of �e is introduced/suggested to explain x-ray luminosities by ic up-scattered photons (e.g., chevalier & fransson, 2006), but indeed i suggest here that one has to check if such a situation is consistent with the radio (cm) properties. for example, for sn 2011dh α ∼> 30 (e.g., �e ∼< 0.3 and �b ∼ 0.01) has been suggested (e.g., soderberg et al., 2012), but as shown above this should produce a detectable change in the radio light curves that was not observed (maeda, 2012). applying the same constraint to a few other sne, it seems like that small �e is a generic feature in sne (maeda, 2013a). 211 keiichi maeda 10 100 0.1 1 10 (e) 25.0 ghz 10 100 0.1 1 10 (f) 36.0 ghz 10 100 0.1 1 10 (d) 16.0 ghz 10 100 0.1 1 10 (c) 8.4 ghz 10 100 0.1 1 10 (b) 4.9 ghz fl ux d en si ty (m jy ) day 10 100 0.1 1 10 (a) 1.4 ghz 10 100 0.1 1 10 (f) 36.0 ghz 10 100 0.1 1 10 (e) 25.0 ghz 10 100 0.1 1 10 (d) 16.0 ghz 10 100 0.1 1 10 (c) 8.4 ghz 10 100 0.1 1 10 (b) 4.9 ghz 10 100 0.1 1 10 (a) 1.4 ghz fl ux d en si ty (m jy ) figure 2: left: multi frequency radio light curves (red solid) as compared with those of sn 2011dh (maeda, 2012). the parameters are (a∗,�e,�b) = (4, 6×10−3, 5×10−2) (left; adopting model a) and (20, 0.26, 2.5×10−4) (right; model b). the synthetic light curves without the ic cooling are also shown (blue dashed). observational data are taken from soderberg et al. (2012). 4 injection and acceleration mechanisms since one can obtain both the spectral and temporal information for sne, there is essentially no degeneracy in deriving the electrons’ injected spectrum slope, p (tab. 1). one interesting issue is found from such analyses – p ∼ 3 is generally derived for young sne, unlike more evolved snrs (mostly p ∼ 2 − 2.4; e.g., bamba et al., 2003) and the standard dsa prediction in the test particle limit (p ∼ 2; e.g., ellison et al., 2000). a cause of the difference has not been clarified, and i propose that this is mainly due to totally different energies of the electrons emitting at cm wavelengths in young sne and more evolved snrs. the argument here is based on that of maeda (2013b). i note that a main difference between the synchrotron emission from sne and that from snrs is that the emitting electrons’ energy is quite different for given frequency (figure 3). typical magnetic field strength is b ∼ 1g for sne (e.g., chevalier & fransson, 2006) and 100µg for snrs (e.g., bamba, et al., 2003). this is consistent with the equipartition expectation (maeda, 2013b). at the observed frequency of ∼ 1 ghz, the emitting electrons’ energies are ∼ 50 mev and 5 gev in sne and snrs, respectively. one can estimate if these electrons satisfy an essential condition required for dsa, namely the electron’s mean free path is exceeding the shock wave width. this is satisfied by electrons with the energy ∼> 100 mev in sne and 10 mev in snrs. thus i suggest that the electrons emitting at ghz frequency are likely in the efficient dsa limit in snrs, while they cannot be efficiently accelerated by dsa in sne. 100 101 102 103 104 105 0.1 1 10 100 1000 dsa (sn) dsa (snr) alma vla /g h z figure 3: the relation between the electron’s energy and the synchrotron frequency, for b ∼ 1g typical of young sne (red-thick-solid) and ∼ 100µg typical of snrs (black-thin-solid). also shown is the minimum electron energy for the efficient dsa, adopting v ∼ 0.1c (sne; red-thick-dashed) and 0.01c (snrs; black-thindashed). the typical frequency coverage is shown by the shaded areas, for cm (‘vla’) and mm (‘alma’) observations . a unified scenario is proposed here – the steep energy spectrum of the electrons derived for young sne reflects the inefficient dsa acceleration, or in other word, the ‘injection’ spectrum. this scenario makes young sne interesting objects in studying the electron injec212 electron acceleration in supernovae and millimeter perspectives tion and acceleration mechanism, as one could directly probe the electron injection mechanism. 5 perspectives for mm observations i propose that observations of nearby young sne at mm wavelengths can potentially provide major advances in the issues discussed in this paper (see maeda, 2013b for details). on the acceleration efficiency, the ic cooling effect is more important at higher frequencies, and thus at mm wavelengths one should be able to see this effect to determine �e, or at least place much stronger upper limit than at cm wavelengths. alternatively, if �e is very small, then the synchrotron cooling becomes important, and the cooling frequency would enter into the mm wavelength. if it happens, it will provide direct estimate of b. in either case, there is a good chance to obtain additional information, and then we can solve the degeneracy between the shock physics and the csm environment (§2). another suggestion is on the electron injection. if the scenario suggested in §4 is correct, we should see the spectral flattening at high frequencies. this flattening could take place already at ∼ 100 mev (§4), then one should be able to detect this signature at mm wavelengths (fig. 3). if such a change in the electrons’ energy spectral slope is detected, this could provide direct evidence of the non-liner acceleration theory where the particles’ spectral slope is expected to become harder for higher energies (e.g., ellison et al., 2004). the energy scale for the possible transition will provide strong constraints on the acceleration theory. for nearby objects (up to ∼ 25mpc), such a signature should be detectable by alma (maeda, 2013b). 6 conclusions in this paper, i have suggested to study electron acceleration mechanisms at a strong shock wave by radio observations of nearby young sne. especially, several ideas have been presented regarding (1) the acceleration efficiency and (2) injection problem and non-linear acceleration toward the efficient dsa. the ideas include (a) to constrain the efficiency by combining radio and optical data, (b) to place an independent constraint on the efficiency by the ic cooling effect, and (c) to probe propertis of ‘injected’ electrons before entering into the efficient dsa regime. i also propose that these issues can be further advanced by mm observations. such observations are being planned – we have our too proposal of nearby sn follow-up observations by alma among the highest priority proposals in alma cycle 1, which is currently active (until early 2014). acknowledgement km thank franco giovannelli and the organizers of frascati workshop 2013 for creating the friendly and stimulating atmosphere. the work by km has been supported by wpi initiative, mext, japan, and by a grant-in-aid for scientific research for young scientists (23740141). references [1] bamba, a., et al.: 2003, apj, 589, 827 doi:10.1086/374687 [2] bell, a.r.: 1978, mnras, 182, 147 [3] bersten, m.c., et al.: 2012, apj, 757, 31 doi:10.1088/0004-637x/757/1/31 [4] blandford, r.d., ostriker, j.p.: 1978, apj, 221, l29 doi:10.1086/182658 [5] björnsson, c.-i., fransson. c.: 2004, apj, 605, 823 [6] chevalier, r.a.: 1982, apj, 258, 790 [7] chavalier, r.a., fransson, c.: 2006, apj, 651, 381 doi:10.1086/507606 [8] ellion, d.c., berezhko, e.g., baring, m.g.: 2000, apj, 540, 292 doi:10.1086/309324 [9] fermi, e.: 1949, phys. rev. 75, 1169 doi:10.1103/physrev.75.1169 [10] fransson, c., lundqvist, p., chevalier, r.a.: 1996, apj, 461, 993 doi:10.1086/177119 [11] maeda, k.: 2012, apj, 758, 81 [12] maeda, k.: 2013a, apj, 762, 14 [13] maeda, k.: 2013b, apj, 762, l24 [14] soderberg, a.m., et al.: 2012, apj, 752, 78 doi:10.1088/0004-637x/752/2/78 discussion sergio colafrancesco: ic losses are dominant w.r.t. synchrotron loses if the b-field is low. how this can match with the expectation that the ic losses are important in high b-field regions in snrs? what is a role of coulomb heating effects? keiichi maeda: on the ic cooling, i believe that we expect that in general the relative importance of the ic cooling is higher for lower b-field. note that i am 213 http://dx.doi.org/10.1086/374687 http://dx.doi.org/10.1088/0004-637x/757/1/31 http://dx.doi.org/10.1086/182658 http://dx.doi.org/10.1086/507606 http://dx.doi.org/10.1086/309324 http://dx.doi.org/10.1103/physrev.75.1169 http://dx.doi.org/10.1086/177119 http://dx.doi.org/10.1088/0004-637x/752/2/78 keiichi maeda talking about cooling, not heating/emission. here, the ic cooling rate is lic ∝ uphγ2 and the synchrotron cooling rate is lsyn ∝ ubγ2. then, for given observed frequency ν, if one increases b then one should decrease γ (to emit at ν), leading to lower lic. in this situation, lsyn can be large b/o the ub term. on the coulomb heating. so far i have been focusing on sne iib/ib/ic, which are believed to have relatively low density csm. i estimated the coulomb effect, and at ghz or higher frequencies, the coulomb heating is estimated to be negligible. 214 introduction non-thermal emissions efficiency of electron acceleration injection and acceleration mechanisms perspectives for mm observations conclusions 161 acta polytechnica ctu proceedings 2(1): 161–164, 2015 161 doi: 10.14311/app.2015.02.0161 study of photometric variability of selected su uma dwarf novae i. voloshina1, v. metlov1 1sternberg astronomical institute, lomonosov moscow state university, moscow, russia corresponding author: vib@sai.msu.ru abstract in this work we present time-resolved photometry of several poorly-studied dwarf novae during recent superoutbursts. observations were made with a ccd, mounted on 50and 60-cm telescopes of the sternberg astronomical institute in crimea in april may and october 2012 and june july 2013. superhumps were detected in light curves of all the dwarf novae. the amplitudes and periods of detected light variations were calculated. superhumps evolution was also followed up for all systems and classification is improved. keywords: cataclysmic variables dwarf novae su uma stars outbursts photometry light curves periods. 1 introduction su uma stars represent one of the three sub-types of dwarf novae. besides the frequent normal outbursts, they could show outbursts of larger amplitudes and much longer duration than the normal ones,superoutbursts. during superoutbursts they exhibit socalled superhumps,an increase of system brightness on the small part of the orbital light curve, that repeats with a period a few percents longer than the orbital one. the amplitudes of superhumps are around 0m.3. the orbital periods of su uma stars are about 80-180 min. there are two small subtypes of dwarf novae inside of this group: wz sge stars with short superhumps period and very long time of outburst recurrence and er uma stars with very short intervals between superoutbursts. it was found from observations that evolution of the superhumps period consists of 3 stages: early evolutionary stage with a longer period of superhumps (stage a), middle stage with systematically varying superhumps period (stage b) and the last stage with a shorter, stable period of superhumps (stage c). all stages are most distinct in wz sge systems. the superhumps, seen during the early stages of wz sge dwarf novae, are double-wave humps and their periods are close to the orbital ones (kato, 2002), the ordinary superhumps are one-wave humps. they have the largest amplitude of periodicity and period slightly longer than the orbital period. according to the tidal-thermal model of osaki (1996) superhumps arise as a result of accretion disk precession triggered by gravitational disturbances from the secondary component. the period of superhumps is beat period between the precession period pprec and orbital period porb. the present paper is aimed at investigating of superhumps phenomenon in selected su uma stars. 2 program stars the stars for our study were chosen from the vsnet list of dwarf novae undergoing superoutbursts during our observational sets. the criteria for our choice were, first of all, the brightness of an object limited by telescope power and relatively rare superoutbursts (and therefore poorly studied). they are: css 121004: 205146-035827 (j2051), v844 her, pnv j19150199+0719471 (j1915), sdss j150240.98+3334239 (j1502). 3 observations the main part of our observations was obtained at the 60-cm telescope with ccd apogee 47 (528 × 512 pxl, pixel size 12×12µm) in v,r and rc bands and partly at the 50-cm telescope with a new ccd apogee alta-8300 (3326×2504 pxl, pixel size 5.4 × 5.4µm) in v band in crimea. the duration of observational sets varied from 3 to 6 h. the accuracy is 2 − 3% for r band, but is less for v band 4 − 5%. the reference stars were taken mostly from aavso list. if we used the local standard star from the vicinity of a program star special high precision observations with ubv photometer at the 60-cm telescope were conducted during the nights with good weather conditions for the standard calibration. ccd data were reduced using the maxim dl standard package, ubv data with software developed by v. lyuty. 161 http://dx.doi.org/10.14311/app.2015.02.0161 i. voloshina, v. metlov 3.1 v844 her the dwarf nova v844 her underwent a superoutburst in may, 2012 (vsnet-alert 14525). this object showed frequent outbursts during 2009–2011, but was relatively inactive next year. the superoutburst that we followed up occurred 370 days after the last superoutburst of 2011. our first observation of v844 her was conducted on april 22 when this nova was in quiescence (∼ 18m) but restarted later after receiving information about the outburst and lasted practically until the end of the outburst when v844 her returned to its pre-outburst brightness. figure 1 show the evolution of superhumps in v844 her. 6052.40 6052.45 6052.50 6052.55 13.1 12.9 12.7 6053.40 6053.45 6053.50 6053.55 13.2 13.0 12.8 6054.40 6054.45 6054.50 6054.55 13.4 13.2 13.0 jd 2450000+ 04.05.2012 05.05.2012 06.05.2012 v v v figure 1: v844 her light curves obtained at the beginning of the outburst. decreasing of superhumps which is clearly seen on the light curve at the bottom indicates that a new stage is approaching. 3.2 sdss j150240.98+3334239 the discovery of outburst of this dwarf nova was reported by vsnet on april 2, 2012. its optical magnitude ranges from 13m.7 to 19m.6 during quiescent and outburst phases, respectively. usno photometry reveals a deep (2.5 mag) eclipse occurring at a period of 84.24 minutes (szkody et al., 2006) which evidence that it is a high inclination system. our observations of j1502 cover 6 nights from april 23 to june 2, 2012. observational light curves for three of them are presented in figure 2. they show deep eclipses besides superhumps. our analysis was restricted by consideration of superhumps phenomenon in this system. figure 2: light curves of high inclination system j1502 showing both superhumps and eclipses. 3.3 css 121004: 205146-035827 the object was discovered by crts on october 4, 2012. observations provided by dr. maehara on october 8 (vsnet-outburst 14685) demonstrate superhumps with figure 3: stage b superhumps in the light curves of dwarf nova j2051. an amplitude of 0m.2 and with a short period of 0d.056 in the light curves of this object. we started our observations later, on october 10 because of the bad sky conditions before this date. at this moment the brightness of j2051 was 14m.9. the object started to fade rapidly on october 20 with rate 0.37m · d−1 and became too faint for observations on october 22 (∼ 18m). our observations of dwarf nova j2051 in v band are shown in figure 3. 162 study of photometric variability of selected su uma dwarf novae 3.4 pnv j19150199+0719471 on june 1, 2013 e.de miguel (vsnet-outburst 15466) reported discovery of an outburst of this transient with magnitude ∼ 10m.8. the next day he reported about detection of early superhumps in the light curve of this object with a possible period of 0d.0641(2) (probably stage a). on june 9 the double-peaked superhumps with an average amplitude of 0m.1 and psh ∼ 0d.0569, were detected by several observers. spectroscopic observations provided by echevarria et al.(2013) suggested the high inclination and low massive component,brown dwarf. our observations of this nova started on june 4, 2013 and lasted until the middle of july 2013. during our observations the brightness of j1915 dropped from 11m.47 at the beginning to 18m (v) at the end of the superoutburst. the light curves of j1915 obtained at the different moments of 2013 superoutburst as an example are shown in figure 4. 6449.3 6449.4 6449.5 11.8 11.7 11.6 11.5 11.4 r c a 6457.3 6457.4 6457.5 12.5 12.4 12.3 12.2 r c b 6467.3 6467.4 6467.5 13.0 12.9 12.8 12.7 r c c 6482.3 6482.4 6482.5 16.0 15.9 15.8 15.7 r c jd2450000 + figure 4: evolution of superhumps in dwarf nova j1915: a beginning of stage a superhumps; b stage b superhumps; c transition from b to c stage; d post-fading stage superhumps. scale is the same for all panels. 4 results for data analysis and search of periodicities we used a special code period98 (sperl, 1998) and period04 (lenz & breger, 2004). the analysis was firstly done for every observational night and later the mean values for different stages of each star were calculated. for v844 her although stages a-c were observed, the period of stage a superhumps was not determined, due to limited observations during this stage. pdot for stage b was positive, as seen in other superoutbursts of this object. in the light curves of dwarf nova j2051 the stages b and c can be distinguished. the superhumps period variations suggests that the object more resemble ordinary su uma-dwarf novae than wz sge-stars. the numerous observations obtained for j1915 (about 6000 images at all) permitted us to follow up superhumps evolution in detail beginning from the first night of our observations, june 4. no superhumps were detected in the light curve of j1915 on this particular night. subsequent observations provided on the next night, june 5, recorded a periodic variation in the light curve of j1915. we suggest that it could be beginning of the stage a. the superhumps period for different stages was determined. the absence of recorded outburst in the past, large amplitude of present outburst, values of parameters (like �,pdot and other) and the pattern of superhumps period variation are typical for wz sge-stars. so we conclude that the recently discovered dwarf nova j1915 belongs to the wz sge-subtype. some of the power spectra calculated for each star as an example are shown in figures 5–8. figure 5: one of the v844 her power spectra as an example. figure 6: the power spectrum of j1502, calculated for all 6 nights of our observations. 163 i. voloshina, v. metlov figure 7: the power spectrum of j2051. figure 8: the power spectrum of j1915. the periods which were determined in frame of this study are collected in table 1. the orbital periods were taken from the literature. table 1: determined periods. object psh(stage b) psh(stage c) porb v844 her 0.05590 ± 2 0.05587 ± 3 0.054643 j2051 0.05715 ± 5 0.05680 ± 6 — j1502 0.05891 ± 2 — 0.06944 j1915 0.05833 ± 2 0.05800 ± 3∗ 0.06164 * for j1915 the period in the second column represent the post-fading superhumps period. 5 summary the main results of our study of su uma dwarf nova photometric variability can be summarized as follows: • the time-resolved photometry in v and r bands was obtained for several su uma-type dwarf novae whose superoutbursts were observed during the 2012–2013 season: – for v844 her 10 nights, ∼ 1400 images in v and rc bands; – for css 120004: 205146-03582 8 nights, ∼ 1000 images in v and r bands; – for pnvj19150199+0719471 13 nights, about 6000 images both in v and rc bands; – for sdss j150240.98+3334239 6 nights, ∼ 750 images in v band. • the individual light curves of all program stars were constructed on the base of these numerous observations. superhumps were detected in the light curves all of them and their evolution during the outbursts was followed up. • the superhumps periods and amplitudes were determined from our data and mean superhumps periods were obtained for each nova (see table 1). • the amplitude of outburst, phase-averaged profile of superhumps and superhumps period variations for css 121004: 205146-035827 suggest that this object more resemble ordinary su uma-stars than wz sge-dwarf nova. • on the base of its outburst properties, the recently discovered nv j19150199+0719471 was classified as a dwarf nova of wz sge subtype. the obtained results are preliminary and reflect the work in progress. they may be slightly improved by using more sophisticated methods and thorough subsequent analysis. acknowledgement i. voloshina gratefully acknowledges professor f. giovannelli for the invitation to take part in this conference as well as his kind hospitality in palermo. the authors thank a. bykov for helping with observations in crimea in 2012. this work was supported in part by the program of state support of russian federation leading scientific schools through grant 2374.2012.2. references [1] lenz, p., breger, m.: 2005, coast 146, 53l. [2] sperl, m.: 1998, coast 111, 1. [3] szkody, p., et al.: 2006, astron. j. 131, 973. doi:10.1086/499308 [4] osaki, y.: 1996, pasp 108, 39. doi:10.1086/133689 [5] . kato, t., et al.: 2012, pasj 64, 21. [6] kato et al.: 2013, pasj 65, 23. (survey iv) 164 http://dx.doi.org/10.1086/499308 http://dx.doi.org/10.1086/133689 introduction program stars observations v844 her sdss j150240.98+3334239 css 121004: 205146-035827 pnv j19150199+0719471 results summary 86 acta polytechnica ctu proceedings 2(1): 86–89, 2015 86 doi: 10.14311/app.2015.02.0086 ae aquarii: a short review p. j. meintjes1, a. odendaal1, h. van heerden1 1department of physics, university of the free state, p.o. box 339, bloemfontein, 9300, south-africa corresponding author: meintjpj@ufs.ac.za abstract the nova-like variable ae aquarii has been continuously studied since its discovery on photographic plates in 1934. in this short review the peculiar multi-wavelength properties of ae aquarii will be reviewed and explained in context of its evolution from a high mass transfer phase, during which period it could have been a supersoft x-ray source (sss). keywords: cataclysmic variable: ae aquarii radiation mechanisms: non-thermal line: identification techniques: spectroscopic. 1 introduction since its discovery on photographic plates in optical wavelengths (zinner 1938), the rapid variable star ae aquarii has been a source of continuous observational and theoretical study involving a wide range of frequencies from radio to tev gamma-rays. optical photometric studies showed that the system is highly variable, with ∆mv ≤ 2 magnitude outbursts occurring nearly continuously (e.g. henize 1949, patterson 1979, van paradijs, kraakman & van amerongen 1989). distance estimates place the source at ∼ 100 pc (welsh, horne & oke 1993). 2 optical and x-ray pulsations the first photometric pulse-timing study of ae aquarii in the optical band (patterson 1979) revealed steady coherent pulsations in the power spectra at p◦ ≈ 33 s and p1 ≈ 16 s, which were interpreted as the spin period of an obliquely rotating white dwarf and its associated first harmonic, caused by the second pole reflecting from the inner edge of the accretion disc (patterson 1979). this led to its initial classification as a dq herculis type system (patterson 1979). fast optical spectroscopy performed later with the mount wilson 2.5 m coudé spectrograph in the wavelength band 636 682 nm, combined with simultaneous photometry in the same wavelength band (welsh, horne & gomer 1993), revealed that the pulsations originate from the white dwarf. however, it has been pointed out (patterson 1979) that the low pulsed fraction (≤ 1%) of the 33 s oscillation of ae aquarii is peculiarly low for a disc accreting white dwarf. in order to place ae aquarii with its rapidly rotating white dwarf in the dq herculis sub-class of the cataclysmic variables, accreting in the slow rotator limit from a well developed accretion disc, a magnetic surface field of b∗ ≤ 100 kg is required (e.g. patterson 1994). this is significantly lower than the inferred surface field of a disc accreting dq herculistype system, which is b∗ ∼ 106 −107 gauss (patterson 1979). optical polarimetry (cropper 1986, stockman et al. 1992; beskrovnaya et al. 1995) revealed circular polarization ∼ 0.05 − 0.1 % in optical light, which confirmed that the dwarf surface magnetic field is at least of the order b∗ ≥ 106 g (e.g. chanmugam & frank 1987). x-ray observations made by einstein (0.1 4 kev) showed the 33 s oscillation, with no indication of the associated 16 s first harmonic (patterson et al. 1980). this is confirmed by recent chandra x-ray observations (e.g oruru & meintjes 2012), which shows a single coherent pulse at 33 s. in optical wavelengths the 33 s period often disappears during outbursts (e.g patterson 1979; meintjes et al. 1992;1994), while the x-ray pulsed fraction determined with the einstein data seems to show some correlation with increasing count rate according to the relation pf(%) = 45(%/s−1)cr(s−1) in the observed count rate range 0.15-0.5 s−1 (e.g. meintjes & de jager 2000). no noticeable changes in the hardness ratio of the x-ray data has been observed during periods of enhanced activity (e.g. meintjes & de jager 2000; oruru & meintjes 2012). the weak correlation between the optical pulsed fraction of the coherent 33 s period and outbursts, together with the radial velocity measurements of uv lines observed during time-resolved spectroscopic observations made with the hubble space telescope hst, suggest a very effective propeller driven mass outflow in ae aquarii (e.g. patterson 1994; eracleous & horne 1996). 86 http://dx.doi.org/10.14311/app.2015.02.0086 ae aquarii: a short review 3 the propeller phase doppler tomography profiles of the ae aquarii system (e.g. wynn, king & horne 1997; ikhsanov, neustroev & beskrovnaya 2004) show tell-tale signatures of a propeller driven mass outflow from the system. the magnetospheric propeller process in ae aquarii was first explained within the framework of large diamagnetic blobs being propelled by a magnetospheric drag (wynn & king 1995; wynn, king & horne 1997). roche tomography of the secondary star (dunford & smith 2005; watson et al. 2006; smith, dunford & watson 2012) shows that ∼ 20% of the surface of the secondary star is covered with starspots, which may suggest a magnetically active secondary star. the effect of secondary star magnetic fields on the mass transfer process has been investigated (meintjes 2004). it has been shown that magnetic prominences can fragment the mass transfer flow into a blob-like stream, which upon interacting with the fast rotating white dwarf field, can be propelled from the system on time-scales that are short compared to the keplerian periods at the distance of closest approach of the stream (meintjes & venter 2005). recent spectroscopy performed with the 1.9 m telescope at sutherland revealed broad emission lines, which is most probably an indicator of high velocity gas in the system. the full-width half maximum width of these lines all implies velocities in excess of vesc ≥ 1000 km s−1 (meintjes, oruru & odendaal 2012), which is of the order of the escape velocity from the radius of closest approach of the stream. it has been shown (venter & meintjes 2006) that the interaction between the fast rotating magnetosphere and the stream of material from the secondary star may result in the development of kelvin-helmholtz instabilities, resulting in effective turbulence driven mixing of magnetic field with plasma. this results in a very effective transfer of angular momentum to the stream, propelling it from the system (venter & meintjes 2006). this has been confirmed by numerical simulations of the interaction between the fast rotating magnetic field of the white dwarf in ae aquarii and the mass flow from the secondary star (bisikalo & zhilkin 2012). the low accretion rate of the white dwarf in ae aquarii (ṁ∗ ∼ 1014 g s−1) compared to the mass transfer rate deduced from the uv emission line spectra (ṁ2 ∼ 1017 g s−1) implies that the white dwarf in ae aquarii is in a super-propeller (ejector) state (bisikalo & zhilkin 2012; ikhsanov & beskrovnaya 2012). the brightening of the source observed in optical on a regular basis, can therefore not readily be explained in terms of enhanced mass accretion onto the white dwarf. it has been shown (beardmore & osborne 1997) that the flaring activity in ae aquarii can be explained satisfactorily within the framework of colliding blobs propelled form the system by the magnetospheric propeller. these authors showed that collisions between faster and slower blobs will result in heating and resultant radiative cooling which can account for the outbursts in ae aquarii. 4 spin-down and particle acceleration the dissipated mhd power due to the magnetospheric propeller process in ae aquarii (e.g. meintjes & de jager 2000; meintjes & venter 2005) is of the order of pmag ∼ 1034 erg s−1, and comes at the expense of the rotational kinetic energy of the rapidly rotating white dwarf. a detailed pulse timing study of the white dwarf spin period, using a 14 year baseline (de jager et al. 1994) revealed that the white dwarf is spinning down at a rate of ṗ∗ ∼ 5.64 × 10−14 s s−1. this translates to a spin-down luminosity of the order of ls−d = iωω̇ ∼ 1034 erg s−1. a more recent study (mauche 2006) revealed that the spin-down rate of the white dwarf has sped-up. this can possibly be explained within the framework of mass transfer variations from the secondary star. the low accretion in comparison to the mass transfer α = ṁ∗/ṁ2 ∼ 0.1% may indicate that the white dwarf in ae aquarii is currently in a super-propeller (bisikalo & zhilkin 2012) or ejector phase (ikhsanov & beskrovnaya 2012) and that the white dwarf in ae aquarii may exhibit the same properties as a spun-up radio pulsar (e.g ikhsanov & bierman 2006). this implies that the low mass accretion and rapid rotation of the highly magnetized white dwarf may provide a mechanism for pulsar-like particle acceleration and nonthermal emission. recent results from suzaku x-ray satellite (terada et al. 2008) showed a non-thermal spectrum at energies �x ≥ 10 kev. it has been showed recently (e.g. oruru & meintjes 2012) that particles can be accelerated in the magnetosphere of the white dwarf to energies in excess of � ≥ 100 gev, which may radiate synchrotron radiation at energies of the order of �x ∼ 10 kev in the weak magnetic field outside the corotation radius. 5 non-thermal emission radio observations performed in the late 1980’s showed that ae aquarii is a non-thermal emitter (bookbinder & lamb 1987, bastian, dulk & chanmugam 1988), showing continuous radio outbursts with maximum flux smjy ≤ 15 at frequencies ν ≤ 22.5 ghz. the high brightness temperature measured tb ≥ 1010 k definitely implies non-thermal emission, with the spectral slope smjy ∝ να, with α ∼ 0.3 − 0.5 (bastian, dulk & chanmugam 1988). the spectral properties of the radio emission in ae aquarii can be explained in 87 p. j. meintjes, a. odendaal, h. van heerden terms of a superposition of synchrotron emitting plasmoids, which expand and cool radiatively through synchrotron radiation (van der laan 1963; 1966). subsequent studies in infrared with iso (abada-simon et al. 2005) and spitzer (dubus et al. 2007) showed that the smjy ∝ να (α ∼ 0.5) non-thermal spectrum extends to a frequency ν ∼ 2000 ghz (dubus et al. 2007). the non-thermal radio-ir spectrum has been modeled successfully (meintjes & venter 2003; venter & meintjes 2006) in terms of synchrotron emission from relativistic electrons in expanding magnetized plasmoids in the propeller outflow. it has been shown that the frequency where the spectrum turns optically thin to radio emission is of the order of νt ≤ 3000 ghz. reports of pulsed burst-like vhe and tev gammaray emission in the 1990’s (bowden et al. 1992; meintjes et al. 1992;1994) sparked interest in ae aquarii as a non-thermal source of high energy emission. subsequent studies with various modern cerenkov detectors could unfortunately not confirm the earlier reports, leaving the vhe-tev status of ae aquarii still in doubt (e.g. lang et al. 1998; sidro et al. 2008; mauche et al. 2012). 6 the evolution the current properties of ae aquarii is consistent with a high mass accretion history (meintjes 2002; schenker et al. 2002). if one consider initial parameters porb,i ∼ 15h, m1,i ∼ 0.6m�, m2,i ∼ 1.6m� (i.e. qi ∼ 2.7) and r2,i ∼ 1.6r�, it can be shown that the initial thermal time-scale (τth ∼ 6.3 × 106(m2/1.6m�) −1 yr) mass transfer could have been of the order ṁ2,i ∼ 2×1019(m2,i/1.6 m�)(τ/τth)−1 g s−1, still below the eddington value of ṁedd ≤ 7 × 1020(m1,i/0.6 m�) −0.8 g s −1. it has been shown that this could have lasted until a critical q-ratio was reached, i.e. qcrit = 0.73 (meintjes 2002) during which time accretion disc torques could have spun-up the white dwarf to periods around ∼ 30 s over a time scale τsu ∼ 3 × 106 yr, which is similar to the thermal mass transfer time scale, i.e. τṁ ∼ few × 106 yr (meintjes 2002). the high mass accretion in this initial phase could have resulted in ae aquarii being a supersoft xray source (sss) (meintjes 2002; schenker et al. 2002). 7 conclusions the multi-frequency properties of ae aquarii have been summarized. the current asynchronicity of the spin and orbital period of ae aquarri can be explained in terms of a high mass accretion history during which period the white dwarf could have been spun-up by accretion disc torques to a short period. in this phase the high accretion rate could have resulted in the system being a sss. a subsequent decrease in the mass transfer rate from the secondary drove the system into the propeller state, which may be the driving mechanism behind the multi-frequency emission from the system. acknowledgement the authors thank the organisers for the invitation to present this work at this conference. references [1] abada-simon, m. et al.: 2005, a&a, 433, 1063 [2] bastian, t.s., dulk, g.a. & chanmugam, g.: 1988, apj, 324, 431 doi:10.1086/165906 [3] beardmore, a.p. & osborne, j.p.: 1997, mnras, 290, 145 doi:10.1093/mnras/290.1.145 [4] bisikalo, d.v. & zhilkin, a.g.: 2012 in golden age of cataclysmic variables and related objects, palermo (september 12-17, 2011), memorie s.a.it, vol 83 (2), p. 562 (eds. f. giovanelli & l. sabau-graziati) [5] beskrovnaya, n.g., ikhsanov, n.r., bruch, a. & shakovskoy, n.m.: 1995, in proc. of the cape workshop on magnetic cataclysmic variables, cape town (1995), asp conf. ser. vol. 85, p. 364, astron. soc. pacific, san francisco (eds. d.a.h. buckley & b. warner) [6] bookbinder j.a. & lamb, d.q.: 1987, apj, 323, l131 doi:10.1086/185072 [7] bowden, c.c.g. et al.: 1992, astropart. phys., 1, 47 doi:10.1016/0927-6505(92)90008-n [8] chanmugam, g. & frank, j.: 1987, apj, 320, 746 doi:10.1086/165592 [9] cropper, m.: 1986, mnras, 222, 225 doi:10.1093/mnras/222.2.225 [10] de jager, o.c., meintjes, p.j., o’donoghue, d. & robinson, e.l.: 1994, mnras, 267, 577 doi:10.1093/mnras/267.3.577 [11] dubus, g., taam, r.e., hull, c., watson, d.m. & mauerhan, j.c.: 2007, apj, 663, 516 doi:10.1086/518407 [12] dunford a. & smith r.c.: 2005, proc. of the astrophysics of cataclysmic variables and related objects workshop, strasbourg (france), asp conference series, vol. 330, p. 399, astron. soc. pacific, san francisco (eds. j.-m hameury & j.-p lasota) 88 http://dx.doi.org/10.1086/165906 http://dx.doi.org/10.1093/mnras/290.1.145 http://dx.doi.org/10.1086/185072 http://dx.doi.org/10.1016/0927-6505(92)90008-n http://dx.doi.org/10.1086/165592 http://dx.doi.org/10.1093/mnras/222.2.225 http://dx.doi.org/10.1093/mnras/267.3.577 http://dx.doi.org/10.1086/518407 ae aquarii: a short review [13] eracleous, m, & horne, k.: 1996, apj, 471, 427 doi:10.1086/177979 [14] henize, k.g.: 1949, aj, 54, 89 [15] ikhsanov, n.r., neustroev, v.v. & beskrovnaya, n.g.: 2004, a&a, 421, 1131 [16] ikhsanov, n.r. & biermann, p.l.: 2006, a&a, 445, 305 [17] ikhsanov, n.r. & beskrovnaya, n.g.: 2012, arxiv: 1205.4330v1 [astro-ph.he] (19 may 2012) [18] lang, m.j. et al.: 1998, astropart. phys., 9, 203 doi:10.1016/s0927-6505(98)00020-6 [19] mauche, c.w.: 2006, mnras, 369, 1983 doi:10.1111/j.1365-2966.2006.10447.x [20] mauche, c.w. et al.: 2012, in golden age of cataclysmic variables and related objects, palermo (september 12-17, 2011), memorie s.a.it, vol 83 (2), p. 651 (eds. f. giovanelli & l. sabau-graziati) [21] meintjes, p.j. et al.: 1992, apj, 401, 325 [22] meintjes, p.j. et al.:1994, apj, 434, 292 [23] meintjes, p.j., oruru, b. & odendaal, a.: 2012, in golden age of cataclysmic variables and related objects, palermo (september 12-17, 2011), memorie s.a.it, vol 83 (2), p. 643 (eds. f. giovanelli & l. sabau-graziati) [24] meintjes, p.j. & de jager, o.c.: 2000, mnras, 311, 611 [25] meintjes, p.j.: 2002, mnras, 336, 265 [26] meintjes, p.j.: 2004, mnras, 352, 416 [27] meintjes, p.j. & venter l.a.: 2003, mnras, 341, 891 [28] meintjes p.j. & venter, l.a.: 2005, mnras, 360, 573 [29] oruru, b. & meintjes p.j.: 2012, mnras, 421, 1557 doi:10.1111/j.1365-2966.2012.20410.x [30] patterson, j.: 1979, apj, 234, 978 doi:10.1086/157582 [31] patterson, j., branch, d., chincarini, g. & robinson e.l.: 1980, apj, 240, l133 doi:10.1086/183339 [32] patterson, j.: 1994, pasp, 106, 209 doi:10.1086/133375 [33] schenker, k. et al.: 2002, mnras, 337, 1105 doi:10.1046/j.1365-8711.2002.05999.x [34] sidro, n. et al.: 2008, international cosmic ray conference (icrc), 2, 715 [35] smith, r.c., dunford, a. & watson, c.a.: 2012, in golden age of cataclysmic variables and related objects, palermo (september 12-17, 2011), memorie s.a.it, vol 83 (2), p. 708 (eds. f. giovanelli & l. sabau-graziati) [36] stockman, h.s., schmidt, g.d., berriman g., liebert, j., moore, r.l. & wickramasinghe, d.t.: 1992, apj, 401, 628 [37] terada, y. et al.: 2008, publ. astron. soc. japan, 60, 387 [38] van der laan, h.: 1963, mnras, 126, 535 [39] van der laan, h.: 1966, nature, 211, 1131 [40] van paradijs, j., kraakman, h. & van amerongen, s.: 1989, a&a, 79, 205 [41] venter, l.a. & meintjes, p.j.: 2006, mnras. 366, 557 [42] watson, c.a., dhillon, v.s. & shahbaz, t.: 2006, mnras, 368, 637 [43] welsh, w.f., horne, k., & oke,j.b.: 1993, apj, 406, 229 [44] welsh, w.f., horne, k., & gomer, r.: 1993, apj, 410, l39 [45] wynn, g.a. & king, a.r. 1995:, mnras, 275, 9 [46] wynn, g.a., king, a.r. & horne, k. 1997: mnras, 286, 436 [47] zinner, e.: 1938, astron. nach., 265, 345 discussion dimitri bisikalo: is there any observational evidence of mass flow surrounding the white dwarf? pieter meintjes: there is observational evidence of mass being ejected from the system. the propeller is extremely effective, only ∼ 0.1% of the mass transfer flow gets accreted onto the surface. the rest is ejected from the system. optical and uv spectroscopy indicate high velocity flows. there may be a circumbinary ring of ejected matter present surrounding the system. 89 http://dx.doi.org/10.1086/177979 http://dx.doi.org/10.1016/s0927-6505(98)00020-6 http://dx.doi.org/10.1111/j.1365-2966.2006.10447.x http://dx.doi.org/10.1111/j.1365-2966.2012.20410.x http://dx.doi.org/10.1086/157582 http://dx.doi.org/10.1086/183339 http://dx.doi.org/10.1086/133375 http://dx.doi.org/10.1046/j.1365-8711.2002.05999.x introduction optical and x-ray pulsations the propeller phase spin-down and particle acceleration non-thermal emission the evolution conclusions 188 acta polytechnica ctu proceedings 2(1): 188–191, 2015 188 doi: 10.14311/app.2015.02.0188 sw sex stars, old novae, and the evolution of cataclysmic variables l. schmidtobreick1, c. tappert2 1european southern observatory, casilla 19001, santiago 19, chile 2departamento de f́ısica y astronomı́a, universidad de valparáıso, avda. gran bretaña 1112, valparáıso, chile corresponding author: lschmidt@eso.org abstract the population of cataclysmic variables with orbital periods right above the period gap are dominated by systems with extremely high mass transfer rates, the so-called sw sextantis stars. on the other hand, some old novae in this period range which are expected to show high mass transfer rate instead show photometric and/or spectroscopic resemblance to low mass transfer systems like dwarf novae. we discuss them as candidates for so-called hibernating systems, cvs that changed their mass transfer behaviour due to a previously experienced nova outburst. this paper is designed to provide input for further research and discussion as the results as such are still very preliminary. keywords: cataclysmic variables dwarf novae novae sw sextantis stars. 1 introduction in the following we briefly review our current knowledge on the populations of old novae and sw sextantis stars especially in the context of the evolution of cataclysmic variable stars (cvs). 1.1 evolution of cvs in general, the evolution of stable interactive binaries is driven by angular momentum loss which as such determines the mass transfer rate. the loss of angular momentum as a source for stable mass transfer implicates that the general evolutionary direction for cvs is from longer to shorter orbital periods. according to the standard model of cv evolution, the main source of angular momentum loss for cvs with orbital periods porb > 3 h is magnetic braking. due to the continuous mass transfer, the secondary is pushed out of thermal equilibrium and becomes bloated. at a period of about 3 h, i.e. at the upper edge of the period gap, the secondary becomes fully convective, magnetic braking ceases and only the much weaker braking through the emission of gravitational radiation continues. therefore, the mass transfer is greatly diminished which allows the secondary star to relax into a state of thermal equilibrium and to contract to a volume corresponding to its mass. it thus loses contact with its roche lobe and mass transfer stops completely. the angular momentum loss via gravitational radiation shrinks the orbit of the now detached and therefore hardly detectable binary until the secondary star fills its roche lobe again at an orbital period of about 2 h, i.e. at the lower edge of the gap and the binary continues as a low mass transfer cv below the gap. for an extensive review of the current understanding of cv evolution, see knigge et al.̃(2011). 1.2 sw sex stars this sub-class was originally defined by thorstensen et al. (1991). incorporating eclipsing nova-like stars with single-peaked emission lines, high velocity line wings, strong he ii emission but no polarisation, and transient absorption features at orbital phases around φ = 0.5. the radial velocity curves show an offset of 0.2 cycle with respect to the phase defined by the eclipse. today, sw sex stars are considered novalike stars with an extremely high mass transfer rate (see e.g. rodŕıguez-gil et al. 2007a,b). this is supported by the high temperatures found for the white dwarfs in these systems (townsley & gänsicke, 2009). while sw sex stars used to be considered as rare and strange objects with unusual behaviour, it has been shown in the last years that they in fact represent the dominant cv population in the orbital period range between about three and four hours (schmidtobreick et al., 2012). from a sample of cvs that was purely seleceted to have an orbital period between three and four hours, they find that the percentage of sw sex stars in this range must exceed 85%. according to the standard model, all long period cvs have to evolve through this period range before entering the gap and will thus share the sw sex characteristics 188 http://dx.doi.org/10.14311/app.2015.02.0188 sw sex stars, old novae, and the evolution of cataclysmic variables during that time. this makes the sw sex phenomenon a phase in the secular evolution of cvs (schmidtobreick et al. in preparation). 1.3 old novae classical novae are cvs that experience a thermonuclear runaway on the surface of the white dwarf where the accreted material has reached a critical mass. in this process, the binary is not destroyed and mass transfer is usually re-established within a few months. it is therefore safe to assume that nova explosions are recurrent events and part of the evolution of every cv with a sufficiently high mass transfer rate to accumulate the necessary material on the white dwarf surface within its lifetime. inbetween the nova events, the behaviour of the cv depends on properties like orbital period, mass-transfer rate, and magnetic field of the white dwarf that also determine its subtype. in addition, the hibernation model predicts changes of the mass transfer rate in the evolution of the preand postnova: after an initial phase of high mass transfer rate – due to the secondary star being driven strongly out of thermal equilibrium via irradiation from the eruptionheated white dwarf – which increases the distance between the two stars, the binary should descend into a long state of low mass transfer once the white dwarf cooled down and the secondary is allowed to relax more towards thermal equilibrium. (shara et al., 1986; prialnik & shara, 1986). some potential evidence for hibernation has been presented in the form of old nova shells around cvs that have previously been known as low mass transfer systems, i.e. z cam (shara et al., 2007) and at cnc (shara et al., 2012). however this interpretation is not exclusive, as the presence of nova shells around dwarf novae could also indicate that all types of cv (including dwarf novae) can experience a nova explosion during their lifetime without necessarily undergoing cyclic changes of the mass transfer rate. 2 what can we learn combining our knowledge from old novae and sw sex stars? several years ago, we conducted a project investigating old novae which had experienced large outburst amplitudes (schmidtobreick et al., 2003; 2005). the idea behind this was that since the absolute magnitude of a nova explosion depends mainly on the mass of the white dwarf (livio, 1992) it thus differs only slightly for different systems. thus, novae with large outburst amplitudes are likely to be intrinsically faint cvs and therefore candidates for low mass transfer rate systems. two systems of our sample, v842 cen and xx tau, show spectroscopic properties that, compared to other old novae, indicate a rather low mass transfer rate. the spectra present comparatively strong balmer emission lines from hα bluewards down to h11. a weak heii emission line is present at 469 nm and indicates a hot component in the system, but at the same time the presence and strength of the hei series is evidence for a significant amount of cooler material. we note that most old novae do not share these characteristics (e.g., ringwald et al. 1996, tappert et al. 2012, 2014). for xx tau, preliminary analysis of long-term photometric monitoring covering a range of 150 d furthermore shows a probable dwarf-nova like outburst with an amplitude of ∼0.8 mag and a duration of ∼10 d (tappert, private communication), which represents additional evidence for a comparatively low mass transfer rate. the approach by tappert et al. (2012) to re-discover lost old novae revealed two more such candidates from spectroscopy: v2109 oph and v728 sco. in particular v728 sco which has also been observed photometrically can be considered a low mass transfer system as it also shows frequent stunted dwarf novae outbursts (tappert et al., 2013a). for three of these low mass transfer old novae, i.e. v842 cen, v728 sco, and xx tau, at least rough values for the orbital period could be established (woudt et al., 2009; luna et al., 2012; tappert et al., 2013a; rodŕıguez-gil & torres, 2005) and they all fall into the range of sw sex stars, i.e. between 3 and 4 hours. on the one hand, this period range is thus dominated by high mass transfer cvs. on the other hand, also the majority of old novae have high mass transfer rates. in fact, the period distribution of old novae shows a distinctive peak in the 2.8-5 h period range, while few novae are found in the period bins that are dominated by dwarf novae (e.g., tappert et al. 2013b, schmidtobreick & tappert 2014). thus, the indicated low mass transfer nature of v842 cen, v728 sco and xx tau is somewhat surprising. for the sake of completeness, we point out that other low mass transfer novae are known. for example, the system v446 her (porb = 4.97 h, thorstensen & talor, 2000) shows frequent ”stunted” dwarf-nova like outbursts (honeycutt et al. 1998). however, we here restrict our discussion to systems with orbital periods in the 3-4 h range. when thinking about a reason why several old novae would appear as low mass transfer systems even though some of them are even situated in the sw sex regime, hibernation comes naturally to mind. it seems an attractive explanation that these cvs have indeed been sw sex stars which due to the nova explosion in the recent past were pushed into a low state and thus experience low mass transfer rates at the moment. in 189 l. schmidtobreick, c. tappert this context we point out that v728 sco in outburst presents a triangular eclipse shape that is typical for sw sex stars (tappert et al., this volume). 3 conclusions and future work cvs with orbital periods between 3 and 4 hours are in general sw sex stars and as such experience high masstransfer rates. however, we find evidence for three old novae with orbital periods in this range that should be of sw sex type but are instead low mass transfer systems. we tentatively conclude that the mass transfer of these systems was changed by the nova eruption, similar to what is proposed in the hibernation scenario. whether the mass transfer in these systems will completely cease or just remain on a low level before rising up again remains to be investigated. to better understand the recent mass transfer history and to thus shed light on the evolutionary state of these systems, it would be important to measure the dynamical masses and the temperatures of the two components in order to determine the binary solution. for this, it is valuable to have a system like v728 sco in the sample which as eclipsing binary offers more opportunities to determine the binary parameters. to test the idea that the low mass transfer novae in the sw sex regime are affected by hibernation, it will be essential to check for the few existing dwarf novae in the sw sex range whether they have experienced a nova eruption in the past. we thus started an observational project to look for nova shells around these objects. other possible candidates for hibernating cv could be found among the so-called pre-cvs, white-dwarf/mainsequence binary which are detached systems. examples for such candidates are ltt 560 (tappert et al., 2011) and qs vir (drake et al., 2014, and references therein) which both have orbital periods in the sw sex range and show evidence of accretion. if indeed a significant number of these low mass transfer systems in the sw sex regime can be shown to have experienced a nova outburst in the past, this would not only be strong evidence for the hibernation scenario but at the same time also prove the evolutionary significance of the sw sex stars as it would yield a natural explanation for the few remaining cvs in the 3-4 h period range that do not follow the sw sex behaviour. acknowledgement this research was supported by fondecyt regular grant 1120338 (ct). we gratefully acknowledge the use of the simbad database, operated at cds, strasbourg, france, and of nasa’s astrophysics data system bibliographic services. we thank an anonymous referee for valuable comments. references [1] knigge c., baraffe i., patterson j., 2011, apjs 194, 28 doi:10.1088/0067-0049/194/2/28 [2] drake j. j., garraffo c., takei d, gänsicke b., 2014, mnras 437, 3842 doi:10.1093/mnras/stt2186 [3] honeycutt r. k., robertson j. w., turner g. w., henden a. a., 1998, apj 495, 933 doi:10.1086/305299 [4] livio, m. 1992, in vina del mar workshop on cataclysmic variable stars, edited by n. vogt, asp conf. ser. 29, 4. [5] luna g. j. m., diaz m. p., brickhouse n. s., moraes m., 2012, mnras, 423, l75. doi:10.1111/j.1745-3933.2012.01260.x [6] prialnik, d., & shara, m. m. 1986, apj, 311, 172. doi:10.1086/164763 [7] ringwald f. a., naylor t., mukai k., 1996, mnras 281, 192 doi:10.1093/mnras/281.1.192 [8] rodŕıguez-gil p., schmidtobreick l., gänsicke b. t., 2007a, mnras 374, 1359 doi:10.1111/j.1365-2966.2006.11245.x [9] rodŕıguez-gil p., gänsicke b. t., hagen h.-j. et al., 2007b, mnras 377, 1747 doi:10.1111/j.1365-2966.2007.11743.x [10] rodŕıguez-gil p., torres m. a. p., 2005, a&a, 431, 289 [11] schmidtobreick l., tappert c., saviane i., 2003a, mnras 342, 145 doi:10.1046/j.1365-8711.2003.06523.x [12] schmidtobreick l., tappert c., bianchini a., mennickent r. e. 2003b, a&a, 410, 943. [13] schmidtobreick l., tappert c., bianchini a., mennickent r. e., 2005, a&a, 432, 199. [14] schmidtobreick l., rodŕıguez-gil p., gänsicke b. t., 2012, mem. s.a.it. vol. 83, 610 [15] schmidtobreick l., tappert c., 2014, stella novae: future and past decades. woudt & ribeiro (eds), asp in press [16] shara, m. m., livio, m., moffat, a. f. j., & orio, m. 1986, apj 311, 163 190 http://dx.doi.org/10.1088/0067-0049/194/2/28 http://dx.doi.org/10.1093/mnras/stt2186 http://dx.doi.org/10.1086/305299 http://dx.doi.org/10.1111/j.1745-3933.2012.01260.x http://dx.doi.org/10.1086/164763 http://dx.doi.org/10.1093/mnras/281.1.192 http://dx.doi.org/10.1111/j.1365-2966.2006.11245.x http://dx.doi.org/10.1111/j.1365-2966.2007.11743.x http://dx.doi.org/10.1046/j.1365-8711.2003.06523.x sw sex stars, old novae, and the evolution of cataclysmic variables [17] shara m. m., martin c. d., seibert m., rich r. m., salim s., reitzel d., schiminovich d., deliyannis c. p., sarrazine a. r., kulkarni s. r., ofek e. o., brosch n., lépine s., zurek d., de marco o., jacoby g., 2007, nature 446, 159 [18] shara m. m., mizusawa t., wehinger p., zurek d., martin c. d., neill j. d., forster k., seibert m., 2012, apj 758, 121. [19] tappert c., gänsicke b. t., schmidtobreick l., ribeiro t., 2011, a&a 532, 129 [20] tappert c., ederoclite a., mennickent r. e., schmidtobreick l., vogt n. 2012, mnras, 423, 2476. [21] tappert c., vogt n., schmidtobreick l., ederoclite a., vanderbeke j., 2013a, mnras 431, 92 [22] tappert c., schmidtobreick l., vogt n., ederoclite, a., 2013b, mnras 436, 2412 [23] tappert c., vogt n., della valle m., schmidtobreick l., ederoclite a., 2014, mnras in press [24] thorstensen j. r., taylor c. j., 2000, mnras 312, 629 [25] thorstensen j. r., ringwald f. a., wade r. a., schmidt g. d., norsworthy j. e., 1991, aj 102, 272 [26] townsley d. m., gänsicke b. t., 2009, apj 693, 1007 [27] woudt, p. a., warner, b., osborne, j., page, k. 2009, mnras, 395, 2177. discussion vitaly neustroev: you claim that some old novae have a very small mass transfer rate. does it mean that these stars should show dn-type outbursts? linda schmidtobreick: yes, indeed. and some stars do show such outbursts. we don’t yet have a lot of data but we have observed several outbursts for v728 sco. 191 introduction evolution of cvs swsex stars old novae what can we learn combining our knowledge from old novae and swsex stars? conclusions and future work 111 acta polytechnica ctu proceedings 2(1): 111–115, 2015 111 doi: 10.14311/app.2015.02.0111 flickering in cvs and accretion disc viscosity r. baptista1 1departamento de f́ısica, universidade federal de santa catarina, campus trindade, 88040-900, florianópolis, brazil corresponding author: raybap@gmail.com abstract i review observational constraints on accretion disc viscosity inferred from changes of disc structure with time and from disc flickering distributions. the radial run of the disc viscosity parameter in four cases are presented and discussed. keywords: cataclysmic variables dwarf novae accretion discs optical time-series photometry individual: v2051 ophiuchi, ht cassiopeia, v4140 sagitarii, uu aquarii. 1 context accretion discs are cosmic devices where angular momentum and gravitational energy are extracted from matter by an anomalous, still unknown viscosity mechanism, allowing it to be accreted onto a central star (frank, king & raine 2002). currently, the most promising explanation for the disc viscosity is related to magneto-hydrodynamic (mhd) turbulence in the differentially rotating disc gas (balbus & hawley 1991). from the observational standing point of view, because the properties of steady-state discs are largely independent of viscosity, one must turn to observations of timedependent disc behavior in order to obtain quantitative information about disc viscosity. here we adopt the prescription of shakura & sunyaev (1973) for the accretion disc viscosity, ν = αss cs h, where αss is the nondimensional viscosity parameter, cs is the local sound speed and h is the disc scaleheight. non-magnetic1 cataclysmic variables (cvs) are excellent sites for studies of disc viscosity, because of their well constrained binary environment and because their accretion discs are usually the dominant light source at uv and optical wavelengths. in particular, the subclass of dwarf novae (dns) show recurrent outbursts in which their discs brighten by factors 20-100 as a consequence of mass and angular momentum redistribution on timescales of a few days. dn outbursts are explained in terms of either a thermal-viscous disc-instability (dim, lasota 2001) or a mass-transfer instability (mtim, bath 1975). dim predicts matter accumulates in a low viscosity disc (αcool ∼ 10−2) during quiescence, whereas in mtim the disc viscosity is always high (α ∼ 10−1). therefore, measuring α of a quiescent disc is key to infer which model is at work in a given dn. 2 flickering in cvs: a short review flickering are the intrinsic brightness fluctuations on timescales of seconds to dozens of minutes seen in light curves of t tau stars, mass-exchanging binaries and active galactic nuclei (bruch 2000 and references therein). the first step in understanding what causes flickering is to find out where is comes from. clues for the location of the flickering sources in cvs come from the analysis of eclipsing systems. in u gem, flickering is stronger at orbital hump maximum and disappears during the eclipse of the bright spot (bs) – where the mass transfer stream hits the outer edge of its accretion disc. these results led to the conclusion that flickering in this dn arises at the stream-disc impact region, because of either unsteady mass inflow (warner & nather 1971) or post-shock turbulence (shu 1976). on the other hand, the flickering in ht cas disappears during eclipse of the central source and recovers well before the egress of its bs (patterson 1981), indicating that it originates in the inner disc regions or at the wd-disc boundary layer (bl). possible explanations for this disc+bl flickering include mhd turbulence + convection (geertsema & achterberg 1992), unsteady wd accretion (bruch 1992) or events of magnetic reconnection at the disc atmosphere (kawagushi et al. 2000). a series of extensive optical flickering studies along the 90’s (bruch 1992, 1996, 2000) strengthened the idea that there are two different sources of flickering in cvs: in objects with strong anisotropic emission from the bs the bs-stream flickering component dominates, whereas the disc-bl flickering component prevails in objects where emission is mostly from the accretion disc. 1white dwarf (wd) surface magnetic fields bwd < 10 5 g. 111 http://dx.doi.org/10.14311/app.2015.02.0111 r. baptista the power density spectra (pds) of flickering sources show a characteristic power-law behaviour at higher frequencies (∝ f−n), with indexes ranging n = 1 − 2, and a flat slope below a cut-off frequency, fc (bruch 1992). the values of n and fc vary from one object to the other and for the same object at different brightness levels (fig. 1). figure 1: optical (b-band) power density spectra for v2051 oph (top) and ht cas (bottom) at two different brightness levels in their quiescent states (differences between bright, intermediate, and faint are 1-2 orders of magnitude lower than differences between outburst and quiescence). dashed lines show the best-fit powerlaw in each case; the corresponding power-law index n is depicted in each panel. 3 estimating the αss parameter 3.1 time changes in disc structure by measuring the (viscous) timescale tv = r/vr ' r2/ν ' r2/(αss cs h) with which the disc responds to changes in mass input rate (frank et al. 2002), one might infer a spatially-averaged disc viscosity parameter αss, αss ' r cs tv [ h r ]−1 (1) where vr is the viscous radial drift speed. application of this time-lapse mapping technique to accretion discs of dn in outburst lead to αss ' 10−1 (e.g., baptista & catalán 2001). 3.2 flickering relative amplitude geertsema & achterberg (1992) investigated the effects of mhd turbulence in an accretion disc. they found that the convection of turbulent eddies lead to large fluctuations in the energy dissipation rate per unit area at the disc surface d(r), which could be a source of flickering in cvs and x-ray binaries. encouragingly, the pds of these fluctuations resemble those of flickering sources, with a power-law dependency of similar index range and the flat slope at low-frequencies. perhaps more important, their model gives a direct relation between the energy dissipation rate fluctuations and the disc viscosity parameter, providing an interesting observational way to infer the local accretion disc viscosity – by measuring the relative amplitude of the energy dissipation rate/flickering. in this model, the number of turbulent eddies that contribute to the local fluctuation is, n(r) = 4 π r h ( h l )2 , (2) where l is the size of the largest turbulent eddies. the local rms value of the fluctuations σd(r) in the average energy dissipation rate per unit area 〈d(r)〉 is given by, σd(r) ' 2.5〈d(r)〉/ √ n(r) , (3) while the disc viscosity parameter can be written as, αss = 3 νt 2 cs h ' 0.9 ( l h )2 , (4) where νt is the local disc viscosity. if the disc-related flickering is caused by mhd turbulence, it is possible to infer αss from the relative flickering amplitude, σd/〈d〉, αss ' 0.23 [ r 50 h ][ σd(r) 0.05〈d(r)〉 ]2 . (5) in the thin disc limit (h � r), disc regions with flickering amplitudes of a few percent already lead to large local αss values (≥ 10−1). in this scenario, cvs with highly viscous accretion discs are expected to show significant disc flickering component. how can we measure σd/〈d〉? from a large, uniform ensemble of light curves of a given cv it is possible to separate the steady-light component, lowand highfrequency flickering amplitudes as a function of binary phase, to derive corresponding maps of surface brightness distributions from their eclipse shapes and, thereafter, to compute the radial run of the relative amplitude of the disc flickering component (flickering mapping, see baptista & bortoletto 2004). 112 flickering in cvs and accretion disc viscosity figure 2: the radial run of the disc viscosity parameter inferred from disc flickering relative amplitude distributions, for three strong flicker dns and for the nova-like variable uu aqr. dots with solid line show average αss values, while dashed lines indicate their 1-σ limits. horizontal dashed lines depict the typical αcool = 10 −2 value expected for quiescent dn discs according to dim. 4 results and discussion baptista et al. (2011) analyzed an extensive data set of optical light curves of ht cas. their observations frame a 2 d transition from a low state (largely reduced mass transfer rate) back to quiescence, allowing the application of both time-lapse and eclipse mapping techniques to estimate αss and to compare the independently derived results. they find that, in the low state, the gas stream hits the disc at the circularization radius rcirc, and the accretion disc has its smallest possible size. the disc fast viscous response to the onset of mass transfer, increasing its brightness and expanding its outer radius at a speed v ' +0.4 kms−1, implies αss ' 0.3 − 0.5. the newly added disc gas reaches the wd at disc centre soon after mass transfer recovery (∼ 1 d), also implying a large disc viscosity parameter, αss ' 0.5. flickering mapping reveal a minor bs-stream flickering component in the outer disc, and a main disc flickering component the amplitude of which rises sharply towards disc centre, leading to a radial dependency αss(r) ∝ r−2, and to large values αss > 10 −1 for r < rcirc (fig. 2) – in agreement with the time-lapse results. a similar analysis was performed for the dn v4140 sgr (baptista et al. 2012). standard eclipse mapping in quiescence indicate that the steady-light is dominated by emission from an extended disc with negligible contribution from the wd, suggesting that efficient accretion through a high-viscosity disc is taking place. flickering maps show an asymmetric source at disc rim (bs-stream flickering) and an extended central source (disc flickering) several times larger in radius than the wd at disc centre. unless the thin disc approximation breaks down, the relative amplitude of the disc flickering leads to large αss’s in the inner disc regions (' 0.15 − 0.3), which decrease with increasing radius. flickering mapping of the dn v2051 oph reveals that the low-frequency flickering arises mainly in an overflowing gas stream and is associated with the mass transfer process. the high-frequency flickering originates in the accretion disc and has a relative amplitude of a few percent, independent of disc radius and brightness level, leading to large αss ' 0.1 − 0.2 at all disc radii (baptista & bortoletto 2004). in uu aqr, optical flickering arises mainly in tidallyinduced spiral shocks in its outer disc (baptista & bortoletto 2008). assuming that the turbulent disc model applies, its disc viscosity parameter increases outwards and reaches αss ∼ 0.5 at the position of the shocks, sug113 r. baptista gesting that they might be an effective source of angular momentum removal of disc gas. since αss increases by two order of magnitude with increasing radius, it is not surprizing that dobrotka et al. (2012) were not able to reproduce the observed pds of uu aqr with a turbulent disc model of constant αss. 5 summary and future steps the picture that emerges from flickering mapping experiments of three dns and of the nova-like variable uu aqr underscores earlier results, indicating that there are mainly two sources of flickering in cvs: the stream-disc impact region in the outer disc and a turbulent inner accretion disc, the relative importance of which varies from system to system. in combination with an mhd turbulent disc model, flickering mapping provides a powerful probe of accretion disc viscosity by allowing one to derive the local magnitude and the radial run of αss from the distribution of the disc flickering relative amplitude. the large αss values found for the three quiescent dns are critical for the outburst mechanism dispute. they are at odds with dim and indicate that the outbursts of a group of dn (strong flickers) are powered by mass-transfer instabilities. two questions related to this result seem to demand further theory development. the first one is why (and how) mass transfer from the donor star in cvs is unstable? the second one is what is responsible for the difference between the low-viscosity discs of quiescent dim-driven dwarf nova and the highviscosity discs of their mtim-driven counterparts? a good first step towards solving this problem might be asking the related question of what is the influence of the wd magnetic field on the radial distribution of αss (e.g., see bisikalo 2014)? an important test of the mhd turbulent disc model still to be done is to perform a flickering mapping experiment on a bona-fide dim-driven dn to check whether it has a low-viscosity accretion disc (as predicted by dim) and if αss is independent of radius (as generally assumed by dim). acknowledgement this work is supported by cnpq/brazil grant 302.443/2008-8. references [1] baptista, r., catalán, m.s.: 2001, mnras, 324, 599 doi:10.1046/j.1365-8711.2001.04320.x [2] baptista, r., bortoletto, a.: 2004, aj, 128, 411 [3] baptista, r., bortoletto, a.: 2008, apj, 676, 1240 doi:10.1086/528706 [4] baptista, r. et al.: 2011, in the physics of accreting binaries, universal academic press (arxiv 1105.1382) [5] baptista, r., borges, b., oliveira, a.: 2012, iau symposium 285, r.e.m. griffin, r.j hanish & r. seaman (eds.), cambridge univ. press, 278 [6] balbus, s.a., hawley, j.f.: 1991, apj, 376, 214 doi:10.1086/170270 [7] bath, g.t.: 1975, mnras, 171, 311 doi:10.1093/mnras/171.2.311 [8] bisikalo, d.: 2014, this proceedings [9] bobinger, a. et al.: 1997, a&a, 327, 1023 [10] bruch, a.: 1992. a&a, 266, 217 [11] bruch, a.: 1996. a&a, 312, 97 [12] bruch, a.: 2000. a&a, 359, 998 [13] dobrotka, a., mineshige, s., casares, j.: 2012, mnras, 420, 2467 doi:10.1111/j.1365-2966.2011.20210.x [14] frank, j., king, a., raine, d.: 2002, accretion power in astrophysics, cambridge univ. press doi:10.1017/cbo9781139164245 [15] geertsema, g.t., achterberg, a.: 1992, a&a, 255, 427 [16] kawagushi, t., et al.: 2000, pasj, 52, l1 [17] lasota, j.p.: 2001, new astronomy review, 45, 449 doi:10.1016/s1387-6473(01)00112-9 [18] patterson, j.: 1981, apjs, 45, 517 doi:10.1086/190723 [19] shakura, n.i., sunyaev, r.a.: 1973, a&a, 24, 337 [20] shu, f.h.: 1976, in structure and evolution of close binary systems, p. eggleton, s. mitton & j. whelan (eds.), dordrecht, 253 [21] warner, b., nather, r.e.: 1971, mnras, 152, 219 doi:10.1093/mnras/152.2.219 discussion john cannizzo: what is the theoretical expression used in determining αss from the speed of the cooling front? raymundo baptista: inferences of αss from time-lapse mapping of outbursting dns are generally 114 http://dx.doi.org/10.1046/j.1365-8711.2001.04320.x http://dx.doi.org/10.1086/528706 http://dx.doi.org/10.1086/170270 http://dx.doi.org/10.1093/mnras/171.2.311 http://dx.doi.org/10.1111/j.1365-2966.2011.20210.x http://dx.doi.org/10.1017/cbo9781139164245 http://dx.doi.org/10.1016/s1387-6473(01)00112-9 http://dx.doi.org/10.1086/190723 http://dx.doi.org/10.1093/mnras/152.2.219 flickering in cvs and accretion disc viscosity based on measurements of the speed of the heating wave during rise, instead of the cooling wave. [note added: bobinger et al. (1997) adopted the expression αcool = vcool/cs to estimate the disc viscosity parameter from their measured speed of the cooling wave vcool. this expression lead to α values smaller than derived from eq. (1) by a factor (h/r).] 115 context flickering in cvs: a short review estimating the ss parameter time changes in disc structure flickering relative amplitude results and discussion summary and future steps 288 acta polytechnica ctu proceedings 1(1): 288–292, 2014 288 doi: 10.14311/app.2014.01.0288 soft x-ray polarimeter: potential instrumentation and observations herman l. marshall1, norbert s. schulz1 1mit kavli institute, cambridge, ma, usa corresponding author: hermanm@space.mit.edu abstract we present an instrument design capable of measuring linear x-ray polarization over a broad-band using conventional spectroscopic optics. a set of multilayer-coated flats reflects the dispersed x-rays to the instrument detectors. the intensity variation with position angle is measured to determine three stokes parameters: i, q, and u – all as a function of energy. by laterally grading the multilayer optics and matching the dispersion of the gratings, one may take advantage of high multilayer reflectivities and achieve modulation factors >90% over the entire 0.2 to 0.8 kev band. this instrument could be used in a small suborbital mission or adapted for use in an orbiting satellite to complement measurements at high energies. we present progress on laboratory work to demonstrate the capabilities of key components. keywords: polarimetry x-rays. 1 introduction polarimetry is an inherently multiwavelength phenomenon. in the radio band, the detectors provide linear polarization information but for other bandpasses, special steps are required, so an observer has to choose whether to obtain the additional stokes parameters. polarization studies in the optical and radio bands have been very successful. radio polarization observations of pulsars provided “probably the most important observational inspiration for the polar-cap emission model” (taylor & stinebring1986) developed by radhakrishnan (1969), critical to modeling pulsars and still widely accepted (taylor & stinebring). tinbergen (1996) gives many examples in optical astronomy such as: revealing the geometry and dynamics of stellar winds, jets, and disks; determining binary orbit inclinations to measure stellar masses; discovering strong magnetic fields in white dwarfs and measuring the fields of normal stars; and constraining the composition and structure of interstellar grains. perhaps the most important contribution of optical polarimetry led antonucci & miller (1985) to develop the seminal “unified model” of seyfert galaxies, a subset of active galactic nuclei (agn). their paper has been cited in over 1000 papers in 30 years, over 5% of all papers ever written about agn. thus, the extra information from polarimetric observations has provided a fundamental contribution to the understanding of agn. although tens of thousands of x-ray sources are known from the rosat all-sky survey, polarization studies were carried out only in the 1970s and were limited to the brightest sources. in over 40 years, the polarization of only one source has been measured to better than 3σ: the crab nebula (novick et al. 1972; weisskopf et al. 1987). even for bright galactic sources, the polarizations were undetectable or were marginal 2 − 3σ results (silver, et al. 1979; long et al. 1980). furthermore, over the entire history of x-ray astronomy, there has never been a mission or instrument flown that was designed to measure the polarization of soft xrays. because of the lack of observations, there has been very little theoretical work to predict polarization fractions or position angles but there has been some recent progress in support of the now-cancelled mission, gems (the gravitation and extreme magnetism smex; see swank et al. 2010). the x-ray extreme universe satellite (xeus) and the international x-ray observatory (ixo) both had been planned to include a polarimeter but their probable successor, athena+, does not, so there is now no mission that is planned to include a polarimeter. here we describe a few potential scientific studies to be performed with an x-ray polarimetry mission with sensitivity in the 0.1-1.0 kev band that would complement observations with an instrument such as gems. 2 the scientific value of soft x-ray polarimetry there is significant scientific potential of an x-ray polarimeter operating below 1 kev. here, we discuss a few types of potential targets; for details see marshall et al. (2010), schnittman et al. (2013), ghosh et al. (2013). 288 http://dx.doi.org/10.14311/app.2014.01.0288 soft x-ray polarimeter: potential instrumentation and observations figure 1: a prediction of the variation of the polarization percentage (left) and its position angle (right) as a function of energy for agn with varying spin, a/m, and eddington ratio, l/ledd (schnittman & krolik, 2009). such studies were initiated with the prospect of obtaining polarization data in the 2-8 kev band using gems. the figures show that predictions depend strongly on energy. observations with a multilayer-based polarimeter would complement those by an instrument such as gems. 2.1 active galactic nuclei blazars are believed to contain parsec-scale jets with β ≡ v/c approaching 0.995. the x-ray spectra are much steeper than the optical spectra, indicating that the x-rays are produced by the highest energy electrons, accelerated closest to the base of the jet or to shock regions in the jet. brindle (1986) showed that the polarization of blazars increases from the ir to the optical band and can be as high as 25%, indicating that the x-ray polarization should be greater than 30%. the jet and shock models make different predictions regarding the directionality of the magnetic field at x-ray energies: for knots in a laminar jet flow it should lie nearly parallel to the jet axis (marscher 1980), while for shocks it should lie perpendicular (marscher 1985). see also poutanen (1994). mcnamara et al. (2009) recently suggested that x-ray polarization data could be used to deduce the primary emission mechanism at the base, discriminating between synchrotron, self-compton (ssc), and external compton models. their ssc models predict polarizations between 20% and 80%, depending on the uniformity of seed photons and the inclination angle. their spectra are very soft, brightest below 1 kev. theoretical work indicates that agn accretion disks and jets should be 10-20% polarized (mcnamara et al. 2009; schnittmann & krolik 2009, dovčiak, et al. 2011) and that the polarization angle and magnitude should change with energy in a way that depends on the system inclination. schnittmann & krolik (2009) particularly show that the variation of polarization with energy could be used as a probe of the black hole spin and that the polarization position angle would rotate through 90 deg between 1 and 2 kev in some cases, arguing that x-ray polarization measurements are needed both below and above 2 kev (fig. 1). 2.2 isolated neutron stars isolated neutron stars (pavlov & zavlin 1997) and magnetars (heyl, shaviv, & lloyd 2003, lai; & ho 2003) should be bright enough for soft x-ray polarimeters. with temperatures below 0.5 kev, they can only be detected below 1-2 kev. magnetars are thought to be powered by the decay of enormous magnetic fields (1014– 1015 g). these fields are well above the quantum critical magnetic field, where a particle’s cyclotron energy equals its rest mass; i.e. b = m2c3/eh̄ (=4.4×1013 g for electrons). measuring the polarization of the radiation from magnetars in the x-ray band will not only verify the strength of their magnetic fields, but also can provide an estimate of their radii and distances and provide the first demonstration of vacuum birefringence, a predicted but hitherto unobserved qed effect (heyl, shaviv, & lloyd 2003) . the extent of polarization increases with the strength of magnetic field and decreases as the radius increases so compact neutron stars are predicted to be highly polarized, > 80% (heyl et al. 2003) . detailed models of less strongly magnetized neutron star atmospheres show that the polarization fraction would be 10-20% at 0.25 kev averaged over the visible surface of the star (pavlov & zavlin 2000). we can constrain not only the orientation of axes, but also the 289 herman l. marshall, norbert s. schulz m/r ratio for the thermally emitting neutron stars due to gravitational light bending. constraining m/r, impossible from the radio polarization data, is extremely important for elucidating the still poorly known equation of state of the superdense matter in the neutron star interiors. we note that these isolated neutron stars do not produce significant flux above 2 kev, so polarimeters with significant effective area in the 0.1 to 1.0 kev band will be needed to test polarization predictions from neutron star atmospheres. figure 2: bottom: overview of a small telescope designed for soft x-ray spectropolarimetry, based on a design suggested by marshall (2008). a small set of nested mirror shells focusses x-rays through gratings that disperse to an array of detectors about 2 m from the entrance aperture. top left: view from the front aperture. gratings are placed behind the mirror with the grating bars oriented along the average radius to the mirror axis. the gratings are blazed in the directions shown. top right: the view from above the detectors and polarizers. spectra from the gratings are incident on multilayer-coated flats that are tilted along the dispersion axis which contains the zeroth order (dot in center). the angle of the tilt is the same for all mirrors and always redirects the x-rays to the adjacent detector in a clockwise direction. the multilayer coating spacing, d, increases linearly outward from zeroth order, just as the wavelength increases, in order to match the first order wavelength to the peak of the multilayer reflectivity. the detectors face the corresponding mirror. measuring the intensity at a given wavelength as a function of clocking angle then provides i(λ), q(λ), and u(λ). 3 a soft x-ray polarimeter we have been working on soft x-ray polarimetry concepts using multilayer-coated optics for almost 20 years, starting with a very simple design (marshall 1994). briefly, multilayer coatings consist of thin layers of contrasting materials usually one with a high index of refraction and the other with a low value. the input wave is divided at each layer into transmitted and reflected components. when many layers are placed on a surface, then the reflected components may constructively interfere, enhancing the overall reflectivity of the optic. the bragg condition must be satisfied: λ = 2d sin θ, where d = da + db is the thickness of the bilayer consisting of one layer of material a with thickness da, and one layer of material b with thickness db; λ is the wavelength of the incident radiation; and θ is the graze angle. when used at brewster’s angle, θ = 45 deg, the reflectivity of p-polarization is reduced by orders of magnitude, so that nearly 100% of the exiting beam is polarized with the e−vector perpendicular to the plane defined by the incoming and outgoing beams. designs for flight instruments based on multilayer coated optics generallly have limited bandpasses. we now have a new design that overcomes this weakness that we are prototyping in the lab. marshall (2007) 290 soft x-ray polarimeter: potential instrumentation and observations showed that it is possible to develop a laterally graded multilayer-coated mirror that combines with a dispersive optic to obtain a broad bandpass. we have studied how to develop a multilayer polarimeter as a small mission (marshall 2008; marshall et al. 2010). fig. 2 shows the results from a design suitable for a mission of opportunity or a small explorer, using one mirror assembly of a size planned for gems but with a 2 m focal length. this telescope would take 5 min to detect a polarization fraction of 8% for a source like mk 421 at its current intensity. 4 conclusions soft x-ray polarimetry has excellent prospects for scientific study of agn, blazars, and isolated neutron stars. we have designed an instrument capable of detecting a polarization fraction as low as 8% in a suborbital mission. we are prototyping components for such a mission in the lab and will have a flight-like configuration within a few months. acknowledgement this work was funded in part by nasa grant nnx12ah12g. references [1] antonucci, r. r. j., miller, j. s.: 1985, apj, 297, 621. [2] brindle, c., et al.:1986, mnras, 221, 739 doi:10.1093/mnras/221.3.739 [3] dovčiak, m., et al.:2011, apj, 731, 75. doi:10.1088/0004-637x/731/1/75 [4] ghosh, p., et al. 2013, arxiv:1301.5514 [5] heyl, j. s., shaviv, n. j., lloyd, d.: 2003, mnras, 342, 134. doi:10.1046/j.1365-8711.2003.06521.x [6] lai, d., & ho, w. c.: 2003, prl, 91, 071101. doi:10.1103/physrevlett.91.071101 [7] long, k.s., chanan, g.a., and r. novick: 1980, apj 238, 710 doi:10.1086/158027 [8] marscher, a.p.: 1980, apj, 235, 386. [9] marscher, a.p., gear, w.k.: 1985, apj, 298, 114. doi:10.1086/163592 [10] marshall, h.l.: 1994, in proceedings of the spie, s. fineschi (ed.), 2283, p. 75-84. [11] marshall, h.l.: 2007, in proceedings of the spie, s.l. o’dell, g. pareschi (eds.), 6688, p. 66880z. [12] marshall, h.l.: 2008, in proceedings of the spie, s.l. o’dell, g. pareschi (eds.), 7011 [13] marshall, h.l., et al.: 2010, in proceedings of the spie, s.l. o’dell, g. pareschi (eds.), 7732. [14] mcnamara, a.l., kuncic, z., and wu, k.: 2009, mnras, 395, 1507. doi:10.1111/j.1365-2966.2009.14608.x [15] novick, r., et al.: 1972, apjl, 174, l1. [16] pavlov, g.g., zavlin, v.e.: apjl, 490, l91. [17] pavlov, g.g., zavlin, v.e.: 2000, apj, 529, 1011. doi:10.1086/308313 [18] poutanen, j.: 1994, apjs, 92, 607. [19] radhakrishnan, v., et al. 1969, nature, 221, 443. doi:10.1038/221443a0 [20] schnittman, j.d., krolik, j.h.: 2009, apj, 701, 1175. doi:10.1088/0004-637x/701/2/1175 [21] schnittman, j.d., et al. 2013, arxiv:1301.1957. [22] silver,e.h., et al.: 1979, apj, 232, 248. [23] swank, j., et al. 2010, x-ray polarimetry: a new window in astrophysics by ronaldo bellazzini, enrico costa, giorgio matt and gianpiero tagliaferri. cambridge university press, p. 251 doi:10.1017/cbo9780511750809.038 [24] taylor, j.h., stinebring, d.r.: 1986, annreva&a, 24, 285. [25] tinbergen, j.: 1996, astronomical polarimetry, cambridge, uk: cambridge university press, pp. 174. doi:10.1017/cbo9780511525100 [26] weisskopf, m.c., et al.: 1978, apjl, 220, l117. discussion jim beall: this arrangement sounds like it’s pretty delicate. how much of a challenge dos this pose for rocket flight? herman marshall: all but one of the components have been flown on satellite missions. the remaining component, the laterally graded multilayer coated mirrors, are likely to survive based on experience with other ml coated optics, used primarily for solar astronomy. the remaining challenge is to maintain alignment and pointing, needed to match the dispersed x-rays to the ml coated mirrors. the sensitivity 291 http://dx.doi.org/10.1093/mnras/221.3.739 http://dx.doi.org/10.1088/0004-637x/731/1/75 http://dx.doi.org/10.1046/j.1365-8711.2003.06521.x http://dx.doi.org/10.1103/physrevlett.91.071101 http://dx.doi.org/10.1086/158027 http://dx.doi.org/10.1086/163592 http://dx.doi.org/10.1111/j.1365-2966.2009.14608.x http://dx.doi.org/10.1086/308313 http://dx.doi.org/10.1038/221443a0 http://dx.doi.org/10.1088/0004-637x/701/2/1175 http://dx.doi.org/10.1017/cbo9780511750809.038 http://dx.doi.org/10.1017/cbo9780511525100 herman l. marshall, norbert s. schulz to misalignments will require stiff structures while variations due to pointing jitter will require that the ml coatings have a sufficiently broad response at a target wavelength, which ultimately depends on the dispersion of the gratings used. so, mitigating jitter will depend on a tradeoff between improving pointing stability and the availability of high dispersion gratings. 292 introduction the scientific value of soft x-ray polarimetry active galactic nuclei isolated neutron stars a soft x-ray polarimeter conclusions acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0056 acta polytechnica ctu proceedings 4:56–61, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app proton radiography with the pixel detector timepix václav olšanskýa, b, ∗, pavel krista, jiří bílab, carlos granjac, radek maříkd, david chvátila a nuclear physics institute of the czech academy of sciences, husinec-řež, czech republic b faculty of mechanical engineering, czech technical university in prague, czech republic c institute of experimental and applied physics, czech technical university in prague, czech republic d faculty of electrical engineering, czech technical university in prague, czech republic ∗ corresponding author: olsansky@ujf.cas.cz abstract. this article presents the processing of radiographic data acquired using the positionsensitive hybrid semiconductor pixel detector timepix. measurements have been made on thin samples at the medical ion-synchrotron hit [1] in heidelberg (germany) with a 221 mev proton beam. the image contrast is generated by the energy of charged particles imparted after passing through the sample. experimental data from the detector were processed for the establishment of the energy loss of each proton using calibration matrices. the interaction point of the proton on the detector was determined with subpixel resolution by a model fitting of the individual signals in the pixel matrix. three methods have been used for the calculation of these coordinates: hough transformation, 2d gaussian fitting and estimate of the 2d mean. parameters of the calculation accuracy and the calculation time are compared for each method. the final image was created by the method with the best parameters. keywords: radiography, charged particles, timepix, medipix, image processing. 1. introduction radiography is often associated with the x-ray radiography, where the image is formed by an attenuation of the beam after passing through the sample. the image is recorded on a radiologic film, a foil or a memory area detector [2]. the process is similar to conventional photography, where negative is created according to radiation intensity. in contrast to this principle proton radiography can get the imaging information about one pixel with only one particle that has passed through this point [3]. therefore it is necessary to know the energy of the particle. thus the principle of proton radiography is to determine the energy of each particle after passing through the analyzed sample. hence proton radiography is suitable for imaging thin samples. the advantages of proton radiography are high sensitivity, high spatial resolution imaging of thin samples and the imaged objects receiving a lower dose [4]. although the absorbed dose by a single proton is significantly higher than that of a photon, the total radiation dose is reduced due to the need for only a single proton to display a single pixel [5, 6]. the aim is to create a system of automatic image processing of proton radiography. proton radiography, unlike traditional x-ray radiography, offers new possibilities and specific imaging thin samples, which can achieve greater sensitivity and spatial resolution. 2. pixel detector timepix 2.1. hybrid semiconductor detector timepix the timepix detector [7] is a position-sensitive semiconductor pixel detector of the family of medipix detectors developed at cern [8]. timepix, based on the previous device medipix2 [9], timepix provides perpixel energy sensitivity. the detector consists of a radiation sensitive semiconductor sensor bump-bonded to the timepix asic readout chip. the sensor can be of different material (si, cdte, gaas) and thickness (e.g. 300, 700, 1000, 1500 µm). the detector provides an array of 256 × 256 pixels for a total of 65 536 independent channels. size of one pixel is 55 µm and the full sensor size is 14 × 14 mm (1.98 cm2). each pixel is connected with its analog and digital signal chain including amplifier, discriminator, counter and digital integrator [7]. 2.2. timepix principle ionizing particles produce in the sensor a cloud of charge which is collected as in a standard semiconductor diode. the collection of charges undergoes charge sharing (due to diffusion) and results in a signal containing several pixels forming a so-called cluster of pixels [10]. the distribution of charge around the impact point has the character of a gaussian. data are recorded in frames which are saved to individual file frames which are time stamped. all data of events are saved to one frame during the selected time. events 56 http://dx.doi.org/10.14311/ap.2016.4.0056 http://ojs.cvut.cz/ojs/index.php/app vol. 4/2016 proton radiography with the pixel detector timepix figure 1. one frame with cropped selection. the left figure shows one frame of radiographic data. this frame contains 9 suitable events and few unwanted values which are recorded by x-rays. three suitable events and three unwanted values are shown on the right figure, where it is displayed the cropped area from left figure. which have been recorded at next selected time are saved to next frame. the detector pixels can be independently operated in three ways depending on the mode in which the detector works [7]: • medipix mode: each event is counted in this mode. it can be used to count the event rate. • timepix mode: value which is recorded in this mode corresponds to the time of interaction during the frame. • time over threshold (tot) mode: the time registers the amount of charge which is given by the energy deposited. data which are measured by mode tot is necessary to calibrate into energy. calibration matrices are used for this calculation [11]. four calibration matrices are used a, b, c and t. the calibration coefficients are obtained by calibration procedure using discrete-energy x-rays from laboratory x-ray sources on xrf [11–13]. the energy was calculated according to equation x = ta + y − b + √ (b + ta − y )2 + 4ac 2a , where x is the matrix with pixel values in energy in [kev], y is the matrix which contains pixel values recorded in [tot] counts and matrices a, b, c and t contain calibration values. 3. experiment and data processing 3.1. description of experimental setup and radiographic data radiographic data used in this work were measured at the hit ion synchrotron [1] in heidelberg (germany). the sample was irradiated by almost monoenergetic proton beam with energy 221 mev. a sample consisting thin foils which were stacked one over the other and a fly’s wing were used a sample. the sample was irradiated in air. the timepix detector which worked at tot mode was placed behind the sample. protons were registered by the detector after passing through the sample. data from the detector were structured into 34 files. each of these files contained around 1000 frames. the duration of one frame was 0.2 seconds. the value of tot and coordinate of each pixel where the threshold had been exceeded were stored. the average number of protons events per frame was about 10 events per frame. the impact of proton on detector creates a round track with a diameter around 7 pixels as shown in fig. 1. this event has as character so called cluster (see fig. 1 and fig. 2) that can be fitted by a 2dgaussian. the highest tot values are located in the center of the track (fig. 2). the position of impact on the detector can be accurately determined based on analysis and fit of all pixels of the event recorded. sub-pixel resolution can be achieved as well. the total energy deposited to the detector by one proton corresponds to the sum of the energy values of all pixels in the cluster. 3.2. data processing methods and their comparasion because all data that were contained in the files were not suitable for processing, it was necessary to select a suitable event record, to avoid unadvisable events and prepare the data for further processing. 2d filtering was used for eliminating unwanted and false events. nonzero pixels were calibrated using energy calibration matrices. furthermore approximate coordinates of each protons impact have been found. an area of 11 × 11 pixels was cropped around these approximate coordinates. because the diameter of the event record was about 7 pixels, it was probable that each pixel corresponding to this event record was included in this cropped area. then the area was evaluated whether it would be suitable for further processing. if two or 57 v. olšanský, p. krist, j. bíla et al. acta polytechnica ctu proceedings figure 2. registration of a proton event in timepix operated in tot (energy mode). the color and height of bars represents the energy of one pixel of the detector. more event records overlapped at one cropped area, this area was not suitable for further work. it was necessary to accurately determine the location of protons impact for image reconstruction. three methods were used for calculation of the precise location of the impact of the proton: hough transform [14], 2d gaussian fitting [15] and estimation of 2d mean [16]. these methods were mutually compared and the reconstruction of image was performed from the results of the most suitable method. hough transform. the detection of the circle, described by equations (1), was used for assessing the center of the event by hough transformation. the idea was in presumption that each event record had a circular nature. a matrix with data of the one frame was converted into a binary map. all contour points of event record were found by using the edge detector. coordinates of circles points which had the same radius as a radius of event records were calculated from the coordinates of each edge point. eighty points of circle coordinates were calculated for each edge point. the value of one was added at each point, whose coordinates were calculated in this manner. maximum of added values was subsequently calculated for each event. there is the impact point. equations (1) describes the calculation of the circles coordinates with center at each edge points, x(θ) = x0 + r cos θ, y(θ) = y0 + r sin θ, (1) where x0, y0 are coordinates of edge points, θ is an angle from 0 to 2π. step of angle θ calculation was chosen up to 1/80 of the circle. variables x(θ), y(θ) are coordinates of the new circle for one angle θ, r is the radius of the circle [14]. 2d gaussian fitting. gaussian fitting was used for calculating center coordinates the impact of one proton as another way. it was performed for only cut selection around traces of impacts. data were fitted to 2d gaussian with equation f(x,y) = c1 + c2 exp ( − ( x − c3 c4 )2 − ( y − c5 c6 )2) , where c3 a c5 are coordinates of the center of 2d gaussian. it corresponds to the location of a proton impact. coefficient c1 corresponds to dislocation, c2 is amplitude and c4 and c6 corresponds to spread of the blob. estimation of 2d mean. estimate of 2d mean was used as the last method for determination center of events records. the energy values of events records were normalized, so that the total sum of all values was equal to one. first of all the normalized values were added up vertically. then the results of these sums were multiplied by the respective horizontal coordinates. finally these values were added up too. this result was one number only and it was the horizontal center coordinate. the same method was used for the calculation of the vertical center coordinates. 4. results 4.1. comparision of methods methods have been mutually compared. similar accuracy was reached using methods estimating 2d mean and 2d gaussian fitting. calculation of the estimate of 2d mean value was about 30 times faster than the 2d gaussian fitting. calculation using the hough transform amounted to less accuracy, and it was also significantly slower than the calculation of 2d mean. comparing of the average euclidean distance between the locations of particles impacts detected using each method and their standard deviations are shown by using each method in table 1. compared methods avg. eukl. dist. std [px] [px] gaussian fitting vs. estimate of 2d mean 0.107 0.057 hough transform vs. gaussian fitting 0.749 0.302 estimate of 2d mean vs. hough transform 0.729 0.295 table 1. comparison of euclidean distance between the location of protons impacts and its standard deviation. it is evident from table 1 that similarly accurate results were achieved using methods: 2d gaussian fitting and the estimation of 2d mean. mean euclidean distance between of these methods was caused by mutual displacement in one direction. this shift 58 vol. 4/2016 proton radiography with the pixel detector timepix was caused slight displacement of the resulting reconstructed image only. the representation of centers of events records using different methods is shown in fig. 3. histograms which are shown in fig. 4 and 5 were created on the basis of the calculation euclidean distances between the centers coordinates of events records detected using each method. figure 3. centroid fitting for one event recorded. the calculated centroid coordinates are shown in the middle of this figure. data1 shows centered calculated using the 2d gaussian fitting, data2 shows calculated center using the estimation of 2d mean and a data3 shows the detected center using hough transform. figure 4. histogram euclidean distances between centroids computed by gaussian fitting and estimation the mean 2d. 4.2. image reconstruction the final image was created from centers coordinates of each trace of impact whose calculation was described in the previous section. this image was created from the sum of the total energy. the value of the total energy which was given to the detector by a proton was calculated as the sum of all energy values of one event record. this value was assigned figure 5. histogram euclidean distances between centroids computed by gaussian fitting and hough transform. to the point whose coordinates have been detected as the place of protons impact. if one pixel of the final image was located for more events than one, the average value of these energetic records was assigned this point. points without localized event are assigned zero value. these points appeared as holes in final image. values for these points can be improved in the final image by extrapolating values from the surrounding area. just one proton can suffice for displaying a pixel. each proton passed through the sample carries the information about a point of the sample. the final image can be displayed magnified due to the subpixel resolution of located impacts coordinates. the picture created from all the 33 000 frames was refined at double zoom. localized have been around 340 000 events. this corresponds to an average of 1.3 protons per pixel. an example of such image is shown in fig. 6. 5. discussion and conclusion the resulting images, which were reconstructed from data obtained using the methods of 2d gaussian fitting and estimation of 2d median values were essentially identical. to display a single image pixel was sufficient to locate the coordinates of only one proton impact. because here we do not measure the decline of are irradiance detector, but directly the decrease of the energy particles is achieved by lower radiation load samples compared with other radiological imaging methods for example x-ray radiography. experimentally it was found that it was ideal to enlarge final image twice. in this view, it may well calculate the value of the missing pixel, because in the vicinity of missing values there are many pixels designed so as to have one missing value extrapolated. the algorithm was created for a complete data processing. the data were processed automatically starting from 59 v. olšanský, p. krist, j. bíla et al. acta polytechnica ctu proceedings figure 6. final radiogram. proton radiography of a composite sample consisting of an assembled foil array and a fly’s wing. the final image was made from coordinate locations of proton impacts using estimating the 2d mean. one pixel represents 22.5 × 22.5 microns. there was located on the average 1.3 protons per pixel. the records of the detector and the resulting final image. another advantage of proton radiography in addition to lower radiation burden is the availability of application accelerators. for the future it is planned to apply radiography using charged particle accelerators van de graaff hv2500 – ieap ctu [17] and cyclotron u-120m – npi ascr [18], which would also examine pokes energy charged particles after passing the samples. also, the plan is to try and apply electron radiography to microtron mt25 – npi ascr [19, 20]. at the electron radiography should be an option to display an image using electron scattering in the sample. here, the dispersion is measured and determined by using the aperture, which should be placed on the fourier plane between magnetic lenses respectively quadrupoles [21]. it is also planned to develop a new automated data processing systems from new applications and other optimization of automated processing system previously available data. 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[16] t. wang, i. y. h. gu, p. shi. object tracking using incremental 2d-pca learning and ml estimation. in 2007 ieee international conference on acoustics, speech and signal processing – icassp ’07, vol. 1, pp. i–933–i–936. 2007. doi:10.1109/icassp.2007.366062. 60 http://dx.doi.org/10.1016/j.nima.2015.05.054 http://dx.doi.org/10.1109/nssmic.2007.4436610 http://dx.doi.org/10.1109/tns.1979.4329851 http://dx.doi.org/10.1118/1.1690713 http://dx.doi.org/10.1109/tmi.2015.2509175 http://dx.doi.org/http://dx.doi.org/10.1016/j.nima.2007.08.079 http://dx.doi.org/http://dx.doi.org/10.1016/j.nima.2008.03.046 http://dx.doi.org/10.1109/tns.2002.803788 http://dx.doi.org/http://dx.doi.org/10.1016/j.nima.2008.03.091 http://dx.doi.org/http://dx.doi.org/10.1016/j.nima.2010.06.183 http://dx.doi.org/10.1016/0031-3203(81)90009-1 http://www.mathworks.com/matlabcentral/fileexchange/37087-fit-2d-gaussian-function-to-data http://www.mathworks.com/matlabcentral/fileexchange/37087-fit-2d-gaussian-function-to-data http://www.mathworks.com/matlabcentral/fileexchange/37087-fit-2d-gaussian-function-to-data http://dx.doi.org/10.1109/icassp.2007.366062 vol. 4/2016 proton radiography with the pixel detector timepix [17] c. granja, m. solar, s. pospisil, et al. basic and applied research with tunable mono-energetic neutrons at prague van-de-graaff accelerator. to be published in proc. xii int. topical meeting on nuclear applications of accelerators (accapp 2015). [18] department of accelerators. isochronous cyclotron u-120m. http://accs.ujf.cas.cz/. [19] a. belov, d. chavatil, c. simane, m. vognar. electron and gamma bremsstrahlung beams of jinr and ctu microtrons. czechoslovak journal of physics 50(1):385–386, 2000. doi:10.1007/s10582-000-0079-9. [20] c. granja, p. krist, d. chvatil, et al. energy loss and online directional track visualization of fast electrons with the pixel detector timepix. radiation measurements 59:245–261, 2013. doi:http://dx.doi.org/10.1016/j.radmeas.2013.07.006. [21] f. merrill, f. harmon, a. hunt, et al. electron radiography. nuclear instruments and methods in physics research section b: beam interactions with materials and atoms 261(1–2):382–386, 2007. doi:http://dx.doi.org/10.1016/j.nimb.2007.04.127. 61 http://accs.ujf.cas.cz/ http://dx.doi.org/10.1007/s10582-000-0079-9 http://dx.doi.org/http://dx.doi.org/10.1016/j.radmeas.2013.07.006 http://dx.doi.org/http://dx.doi.org/10.1016/j.nimb.2007.04.127 acta polytechnica ctu proceedings 4:56–61, 2016 1 introduction 2 pixel detector timepix 2.1 hybrid semiconductor detector timepix 2.2 timepix principle 3 experiment and data processing 3.1 description of experimental setup and radiographic data 3.2 data processing methods and their comparasion 4 results 4.1 comparision of methods 4.2 image reconstruction 5 discussion and conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0001 acta polytechnica ctu proceedings 5:1–3, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app problems of en 50 128:2011 railway standard tomáš brandejský czech technical university in prague, department of applied informatics in transportation, faculty of transportation, praha, czech republic correspondence: brandejsky@fd.cvut.cz abstract. the second version of railway standard en50128:2011 published by cenelec is used for five years and thus the cenelec reasons about preparation of new version of this standard are coming, because the cenelec defined 10-year cycle of its standards innovation. the paper discusses some observations of problems of this standard application from the position of safety assessor which gives us the contact with developers applying this standard into the real railway products. keywords: railway software applications, assessment, safety, reliability, accessibility, maintainability, development. 1. introduction since the origin, the group of cenelec’s railway stands [1] [2] [3] was declared as standard of all safety related railway applications including not only interlocking systems, but rolling stock and related communication too. the standard en50126 defines its applicability to "the specification and demonstration of rams for all railway applications and at all levels of such an application, as appropriate, from complete railway systems to major systems and to individual and combined subsystems and components within these major systems, including those containing software". the standard en50128 defines its application domain as: "this european standard specifies procedures and technical requirements for the development of programmable electronic systems for use in railway control and protection applications. it is aimed at use in any area where there are safety implications. these may range from the very critical, such as safety signalling to the non-critical, such as management information systems. these systems may be implemented using dedicated microprocessors, programmable logic controllers, multiprocessor distributed systems, larger scale central processor systems or other architectures". and the standard en50129 defines its scope as electronic systems related to safety in application to railway interlocking systems. it means that at least standards en50126 and en50128 must be reasoned in any railway system related to safety. on the beginning of application of these standards safety assessors required application of these standards especially in the domain of interlocking systems and the rest was frequently omitted. after the year 2011, when the second generation of these standards has come into the operation the situation changed and the conformity with these standards is required also by assessors of rolling stock and other railway systems. it brings some problems described below due to different requirements on system functions in situation, when the system malfunction is detected. the main aim of the paper is the discussion of problems in nowadays application of the standard en50128 and offering of ways how to improve it. 2. problems of en50128 application the cenelec’s standard en 50128 "railway applications – communications, signalling and processing systems – software for railway control and protection systems" was defined as part of three standards group [1] [2] [3] applying standard iec 61508 "functional safety of electrical/electronic/programmable electronic safety-related systems" into the specific area of railway systems. the main difference is in different safety assessment scheme, where not only developer, validator and assessor are reasoned but also railway safety authority as next independent body. in the next chapter, particular problems will be discussed: (1.) the last version of the standard is less transparent than original one. the change of the standard structure was asked by cenelec due to unification of all cenelec standard structure. whilst the original standard was divided into 17 chapters and 2 annexes, the 2011 version is organized into 9 chapters and 4 annexes (2 are new). this change causes not only that there are no separate chapters e.g. for software validation and assessment, but software assurance chapter describing testing, verification, validation and assessment is before chapter 7 describing development process! furthermore, the chapter 7 is named "generic software development" now not looking that the term "generic software" in computer science doscon mean "common sense software", as this term is used in the standard, but "a class of software that can be used for a number of different purposes without requiring modification", as it is defined e.g. in [4] [5]. what is worse, the 1 http://dx.doi.org/10.14311/app.2016.5.0001 http://ojs.cvut.cz/ojs/index.php/app tomáš brandejský acta polytechnica ctu proceedings chapter 3.14 of 50128:2011 standard defines generic software meaning some way. thus the topic of the chapter 7 does not give any sense. similarly in the chapter v 3.1.4 of original en50128 introduces the term "component" which is defined extremely vague (and in the en50128:2011 even it is not defined), but especially in the area on object oriented programming the meaning of this term if defined precisely and totally different. because the standard en50128 is used especially by programmers, such drifting of frequent term sense might cause many misunderstandings. because en50128:2011 allows to understand term component also as object, module or even function, the meaning of e.g. chapter 7.5 might be understand totally different by particular people in particular situations. (2.) there is not explicitly declared applicability of the standard to railway vehicles and especially differences in requirements to behaviours of vehicle control systems in the case of malfunction, which is opposite to required behaviours of interlocking control systems in the analogical situation. typically, in the case of control system error vehicle must keep moving but interlocking system can signalize stop to all tracks and perform e.g. restart of the control system. (3.) there is not solved defence against targeted attack (terrorism). defence of mission critical systems against targeted attacks e.g. requires ability of the system to isolate attacked parts and thus to reconfigure system. it is a big problem, because dynamic reconfiguration is listed in table a.3 as technique 14 and it is not recommended non respecting fact that this technique is frequently adopted e.g. in military, cosmic or avionics systems. it is not possible to constraint this problem only to subjects of safe communication and physical barriers. on the other hand, standards of the information security like iso 27001 are rather oriented to managing sensitive company information than to control systems located in open space and communicating across it. (4.) the standard does not solve design of programmable hw, where on interface of sw/hw occur many specific problems which solution is not defined in the actual standard. many of these problems were discussed in [6], like specific problems of design, timing and application structure but there also another problems like problem of dedication of safety related functions to development tools which is related to problems discusses herein in chapter (6.) and which is not solved. design tools transforming e.g. vhdl language into configuration bit stream of the chip does not work deterministically and this fact is hard to accept in safe system design. (5.) the standard en50128 contains extensive appendix b "key software roles and responsibilities" on ten pages and small appendix c "documents control summary", which only recapitulate textual part figure 1. the example of lifecycle hierarchy of complex system. of the standard and do not bring any additional information. it is good to delete such chapters because they only decrease clarity of the standard. (6.) chapter 8 "development of application data or algorithms: systems configured by application data or algorithms" should explain better requirements of systems with different level of configurability. it is hard to compare simple system which is configured by one dimensional table of parameters and system where safety-related functions are described in configuration language, which is interpreted in real time and where part on safety checks is dedicated to development tools. the figure 1 draws possible configuration of different life-cycles cooperating on development of final (sometimes also generic) application based on configurable generic kernel where e.g. life cycle of generic system determines requirement on lc of final application and final application development tools life cycle. (7.) in comparison to standard iec 61508-3 this railway standard does not contain equivalent of iec 61508-3 appendix c explaining which sw property is achieved by which technique. such explanation is useful both for designers and assessors because it simplifies, objectivises and formalizes selection of design techniques in atypical, non-standard situations where groups of design techniques predefined in the table a.3 are inapplicable. typical example of this situation is in this moment application of programmable hw, which is not solved in the standards and which requires a little bit different approach than standard hw and sw. (8.) there is not solved problem of design of systems with safety integrity level greater than 4. growing complexity of railway systems, especially of such systems as ertms, tends to requirement of such sil. it is true, that many calls for higher sils are given by misunderstandings of sil meaning. sil is related to tolerable hazard rate, as it is introduced in the standard en 50129, table a.1. tolerable hazard rate is related to one hour and one function, but many designers of railway systems missrelates it to number of invocations of the function and means 2 vol. 5/2016 problems of en 50 128:2011 railway standard that systems working on higher operation frequency must have higher sil. it is mistake. (9.) there is not well defined that sils are determined to top-level functions, not to whole system and not to any sub-functions. the standard en50129 uses formulation single function without precision of the complexity of the function. many top-level functions are performed by groups of lower level functions. this lack of precise definition tends to many useless discussions between manufactures and assessors. (10.) the standard defines five different sil levels but de facto it specifies only three different cases: sil0, sil1 and 2; and sil 3 and 4. in addition, requirements to sil 1 and 2 systems are inappropriate strict as well as there are requirements to documentation complexity of sil0 systems, especially in comparison to cots systems of the equal sil level. but in fact it would be just opposite – e.g. it is need to do precise testing developed by other company, which design process is not well described, where the documentation is missing, than in the case of well documented one. 3. discussion ten year cycle of cenelec standard upgrade is difficult especially to rolling stock manufacturers. locomotives, carriages and other vehicles have extremely long life-cycle taking any tens of years. there is long design phase including real research, own assessment phase sometimes takes about ten years and also operation is reasoned in horizon of thirty years or more. thus the idea that between start of design and start of operation the required standard will be once or even two times changes is not acceptable. because cenelec does not accept this fact, rolling stock making companies provide "diverse activity" in related cenelec work-groups and prevents significant changes of these standards. on the opposite side, above mentioned problems does not represent significant changes of the standards and their acceptance in the new generation of them might increase their applicability. 4. conclusions the above presented paper summarises problems on the today railway standard en50128:2011, which were observed in certification body of czech technical university in prague – cov fd čvut v praze during certifications of many railway systems serving especially (but not only) in czech republic. references [1] en50126 railway applications – the specification and demonstration of reliability, availability, maintainability and safety (rams). [2] en50128 railway applications – communication, signalling and processing systems – software for railway control and protection systems. [3] en50129 railway applications – communication, signalling and processing systems – safety-related electronic systems for signalling. [4] http://www.teach-ict.com/glossary/g/ genericsoftware.htm. [5] http://www.computingstudents.com/dictionary/ ?word=generic%20software. [6] t. musil. návrh metodiky pro vývoj a verifikaci bezpečných algoritmøu implementovaných v dynamicky rekonfigurovatelných fpga, phd. thesis, 2015. 3 http://www.teach-ict.com/glossary/g/genericsoftware.htm http://www.teach-ict.com/glossary/g/genericsoftware.htm http://www.computingstudents.com/dictionary/?word=generic%20software http://www.computingstudents.com/dictionary/?word=generic%20software acta polytechnica ctu proceedings 5:1–3, 2016 1 introduction 2 problems of en50128 application 3 discussion 4 conclusions references acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0050 acta polytechnica ctu proceedings 4:50–55, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app ion source for a single particle accelerator tomáš matlochaa, b a faculty of nuclear sciences and physical engineering, czech technical university in prague, břehová 7, 115 19 praha 1, czech republic b nuclear physics institute of the cas, 250 68 řež u prahy, czech republic correspondence: matlocha@ujf.cas.cz abstract. this paper proposes a method to obtain very low beam current on the injection side of an particle accelerator. by using a combination of a well known ion source type and a short photo-emissive laser pulse, very low amount of ions can be created. the method could be used in the ion source of various types of the accelerators, where very low beam current is essential. the design is adapted to build a compact internal ion source and the dimensions of the device are adjusted to fit central region of the cyclotron u-120m. the functional parameters of the device are discussed and the amount of the produced ions is estimated. keywords: single particle accelerator, cyclotron ion source, low beam current, laser ion source, pulsed ion source, u-120m, proton radiography. 1. introduction this paper presents a principle of delivering a very low beam current to the injection system of a single stage or low energy accelerator. a control of the ion creation process is a complex task with effect of many parameters. one of the control parameter can be the length of the primary ionizing pulse in the ion source. this length is naturally restricted and usualy can not be arbitrary lowered. when the ionizing pulse is driven using a high voltage discharge, the pulse length with duration about tens of nanosecond can hardly be achieved. from this reason, a primary ionizing plasma in presented ion source is not made using ordinary high voltage discharge, but using a laser pulse. with the present-day laser systems, the high intensity laser pulses with the duration of few ps could be generated and used for photo-emission. electrons from photo-emission then controllably ionize gas and produce ions for the low energy proton or heavier particle accelerator. a usage of the low current beams vary accordingly with the technical field where the beam is applied. the low current beams are required in the radio-biological microbeam study of the living cells [1], in the single proton detectors tests or in the proton radiography. a benefit of using proton beams instead of classical gamma-rays in the radiography, is that the scanned sample receive significantly lower dose for similar image result [2]. with a proper magneto-optic lens system, very high-resolution images could be obtained. usual resolution even with the low energy proton beams is several tens of µm. this resolution of the proton radiography method could be apparently further improved and its limits are far beyond present ranges [3]. very good resolution of the proton radiography [4] could lead to a specific high resolution 3d imaging, where a complete 3d image of the sample could be obtained by two or three laser shots in the proposed ion source. for a very small and sensitive scanned samples, the lowest possible irradiation can be important, so precise control of the number of the created ions is essential. schematic principle of the proton radiography is shown in fig. 1. figure 1. the proton beam lens system with the particle trajectories as colored lines [3]. 2. beam current limitation the idea presented in this paper has grown from the response tests of the silicone detectors and electronics for the planed upgrade of the inner tracking system of the alice detector [5]. the tests were performed on the old cyclotron u-120m, shown in fig. 2. these tests however require very low intensity of the delivered beam current [6]. a minimal beam current which can be delivered comfortably by this cyclotron is ca. 2.5 fa/cm2, which corresponds to ca. 15 600 protons per second per cm2. the beam intensity is usually decreased only after its extraction from the ion source. it means that regardless of the output cyclotron beam current, the amount of ions at the 50 http://dx.doi.org/10.14311/ap.2016.4.0050 http://ojs.cvut.cz/ojs/index.php/app vol. 4/2016 ion source for a single particle accelerator figure 2. cyclotron u-120m. beginning of the acceleration process is almost the same. this factor limits stability of the accelerator and reduces utilization of low current beams. the laser induced photo-emission ion source presented in this paper, can significantly reduce the number of ions entering the machine. 2.1. u-120m pig ion source the ion source of the cyclotron is a movable part and its position can be fine-adjusted with respect to the puller. the extraction window of the ion source has area approx. 2 mm2. the puller has an input window with an area of about 20 mm2. an optimal position of the ion source opposite to the puller guarantees a maximal extraction to loss ratio. in normal operation this extraction ratio is kept as high as possible. when the low current beam is required, this ratio is lowered. this is achieved mainly by changing the position of the ion source extraction window. the ion source is moved slightly to one side, so the extraction window of the ion source does not overlap with the puller input window. subsequently, only a minimal number of ions are able to pass the initial extraction. the rest of the ions are lost on the puller body as can be seen in fig. 3 and 4. although the above mentioned current regulation is very easy, it has some disadvantages. the u-120m cyclotron ion source is a penning ion gauge (pig) type ion source with a cold cathode, operating in continuous regime. in fig. 5 is photo of the ion source oriented with its head on the left side, and in fig. 6 is its head detail. the ion source has a cylindrical plasma chamber parallel to the cyclotron magnetic field of approx. 1 t. the chamber is enclosed from both sides by tantalum cathodes. the cylindrical wall of the ion source is an anode and whole body is made from a tungsten-copper composite. electrons emitted from the cathodes are accelerated by the anode potential. due the strong magnetic field the electrons are trapped by the field lines and are not able to directly reach the anode. the electrons repeatedly oscillates between the cathodes and are slowly drifting to the anode figure 3. central region top view – optimal extraxion for proton regime 37 mev simulated with durycnm4 [7]. figure 4. central region top view – minimal extraxion for proton regime 37 mev simulated with durycnm4 [7]. body. the electrons ionize the gas in the cylinder and create the plasma. the amount of the produced ions can be slightly regulated by changing the plasma current and by regulating the ion source gas presure. for stable operation, the penning ion source needs some minimal plasma current to heat up cathodes and some minimal corresponding gas concentration. in the usual working regime the plasma current is about 1 a and the minimal plasma current for stable operation is several miliamperes. basically in the ion source are two concurrently 51 tomáš matlocha acta polytechnica ctu proceedings figure 5. u-120m pig ion source. figure 6. current u-120m pig ion source. processes which lead to creation of ions. the first process is based on stripping orbital electrons of neutral atoms and generates positive ions. the degree of ionization is done by temperature of the plasma electrons and by a time of interaction. our penning ion source has plasma electron temperature up to approx. 70 ev. it is enough to produce ions with low ionization potentials, e.g. h+, d+, he2+, li2+, c4+, etc. in second process, the negative ions are produced. the second process leads to production of negative ions via the surface and the volume production process. the section 5.1 describes the negative hydrogen ions formation process in more detail. after their creation, both negative and positive ions are simultaneously extracted from the ion source by an alternating rf high voltage field. negative polarity extracts positive ions and oppositely. polarity of the cyclotron magnetic field determines the type of ions which will be accelerated. the second type of ions are simply extinguished on the puller body. 2.2. pulsed pig ion source from above mentioned penning ion source operation principle it follows, that the ions are produced permanently. from that reason the lowest current of extracted ions is limited from below. there were several attempts to switch our penning ion source to a pulsed operation mode. these studies were motivated by other reasons than lowering the output current, but a deeper knowledge of the pig pulse regime capabilities were gained. an ignition process of the pig ion source from the full off state is very long and lasts couple of seconds. when the pig plasma is once ignited, the plasma current can be lowered to five miliamperes, and the ion source is in low current state. in this state changing of the plasma current can be done quickly with a reaction time of a few microseconds. then the ion source can be fast and easily modulated between low and high current states. but the ion source even in the low plasma current mode is still producing much more ions than required. there is an option to further lower the low current value by at least an order of magnitude by driving the cathodes temperature. nevertheless, continuous operation of the ion source is not suitable for the low current beams. from this reason, the possibilities of other ion production were studied and the main findings are presented in the next paragraphs. 3. inspiring ion sources 3.1. ebis/ebit ion source a principle of operation of an electron beam ion source (ebis) and an electron beam ion trap (ebit) is very similar. the electrons are extracted from a cathode by thermionic emission and accelerated in a strong solenoidal magnetic field by an anode voltage. the shape of the magnetic field forms the entering electrons to well focused beam with densities up to several ka/cm2. this beam ionizes the gas and forms the ions. because the electron beam temperature can be very high, a high ion charge state is possible. in the ebit the ions are confined in a space charge of a dense electron beam, where attractive force of electrons prevent the ions to move. an electrostatic potential of external electrodes is applied to trap the ions and hold their position. when the trapping potential is switched off, the ions can be extracted. this type of ion source is mainly used for heavy ions for which high charge states or full stripping by other techniques is complicated. temperature of electrons could be from 10 kev to 200 kev. the extracted ion current is low, typically ranging from 105 to 109 ions per second. the ebis/ebit sources exist in many designs from ultra compact devices for low current and low charge states to huge machines capable of delivering fully stripped uranium ions [8, 9]. principial schematic of the ebit is in fig. 7. figure 7. schematic design of ebit ion source [10]. 52 vol. 4/2016 ion source for a single particle accelerator 3.2. rf gun modern light sources such as the energy recovery linac (erl) or free electron lasers (fel) use as an electron beam generator pulse operated rf guns. the rf gun is a laser induced photo emission electron gun, where photo electrons are accelerated by a strong electric field generated during the rf pulse. schematic design is shown in fig. 8. the photosensitive cathode is placed near the center of a 3/2 cell rf cavity (0.5–2 ghz), where a high electric field is produced (up to 120 mv/m). a very short laser pulse (10 ps) with a proper phase is applied onto the photo-cathode. a bunch of photoelectrons is emitted, accelerated to proper energy and delivered further into the system. currently exist a number of devices of this type, capable to produce electron beams with energies in the range 0.1–5 mev. figure 8. schematic design of the rf gun. photo-cathodes used in these electron sources can be from metal or semiconductor. in general, metals have lower electron conversion efficiency(qe) and need uv lasers. on the other hand they have fast response and are resistant to impurities. semiconductor cathodes are more efficient and can be operated with visible/ir laser, but have slower time response and are sensitive to contamination and photon radiation. charge emitted per bunch is approx. 1 nc. required peak rf power is up to 10 mw. mean laser power is 1 w and required frequency is few thousands pulses per second [11, 12]. 4. laser induced photo-emission ion source the above mentioned devices could probably be combined. by applying a short laser pulse on a photo sensitive cathode one could produce an electron bunch. this bunch could be then accelerated to a desired energy by an anode potential. during travel from a photo-cathode to an anode, this bunch could ionize gas present in the discharge chamber. because the electron temperature is expected only few tens ev, sufficient to ionize hydrogen and helium, and emitted charge is low, the device could be kept compact. such a compact device could be placed inside the central region of the cyclotron and in demand of very low beam currents replace the present pig ion source. dimensions of the pig ion source must be preserved to fit into the cyclotron central region. these dimensions determine the discharge chamber shape and possible arrangement of the electrodes. unlike in the pig ion source, where the cylindrical chamber is anode enclosed by cathodes from both sides as in fig. 6, here will be photo-cathode on one side and anode on the other side of the cylinder as in fig. 9. in this configuration the electron bunch will pass through the chamber only once per one laser pulse. the laser pulses with energies of few µj are driven to the photo-cathode through ion source arms enclosed by a high vacuum sealed window, from an external laser generator placed out of the cyclotron site. figure 9. proposed laser ion source. 4.1. photo-cathode in general, the photo-cathode should have work function below 3 ev to avoid a necessity of using uv lasers. considering a gamma radiation always present at the cyclotron site, the photo-cathode should be metallic. other reason for this choice is good resistance to adverse conditions such as bad vacuum and impurities in the system, where semi-conductors cathodes are very sensitive. in theory, the metal photo-cathodes should not work when irradiated by photons with energies below the work function hν < φ. in reality, an effect of multiple photon absorption were observed in high density laser pulses, and even photons energies below the threshold energy can be used for photoemission [13]. the quantum efficiency is defined by ratio between incident photons and emitted electrons. the metals have qe low, but in our case it should be helpful for limiting a parasitic emission by photons captured from de-excitation and neutralization of the gas atoms. copper, which has qe 5 × 10−5 at wavelength 320 nm, could be suitable for the first try. 53 tomáš matlocha acta polytechnica ctu proceedings 5. ions production 5.1. negative hydrogen ion formation as the principle of proposed pulsed ion source is simple, many practical limitations will probably occur. the main questions are, if this device will be capable to produce any negative ions and how long this ions will live for. a creation of the negative hydrogen ion h– is a multi-step process, where several conditions should be fulfilled. there are many processes involved in h– formation. one of them, the volume negative ion production process which needs at least two different temperature of electrons. faster electron with energies near 20 ev which ro-vibrationally excites the hydrogen h*2 and a cold electron, which can than dissociate the h2 molecule and attach to one of the atoms to form an negative hydrogen ion: h2 + e − −−→ h∗2 + e − −−→ h0 + h−. this process is called a dissociative attachment process and the dissociative low temperature electron energy needs to be close to 1 ev. bounding energy of the second electron in the h– ion is 0.75 ev, so the once formed h– ion can be easily destroyed by collisions with other ion or on residual gas. the second important h– formation process is the surface negative ion production. in this process the previously formed h0 atom from h2 dissociated molecule collides with a metallic surface of the plasma chamber. when bouncing back from the wall, an conductive electron from the metal surface can be trapped by the atom and form an h– ion. work function of the chamber metal tungsten surface is near 4.5 ev, far above 0.75 ev electron affinity of h atom. from this reason the cross-section of h– formation from pure metal surfaces is commonly increased by covering surface with caesium layer. this thin cs layer lowers the work function of surface to approx. 1.6 ev. the cs is widely used in h– ion sources and has its significant role also in the volume h– production. in both processes significantly increases h– formation probability. in the u-120m cyclotron penning ion source cs is not used, so the volume process is therefore considered to be more important [14–16]. 5.2. positive ions formation a situation with positive ions should be more simple. ionization of a low atomic number elements is done by stripping their orbital electrons in energetic collisions. as for heavier elements the process is gradual, because the atoms are ionized in a stepwise processes where each electron collision removes one or two orbital electrons. for light atoms such as h and he, this can be done very fast. ionization energy for hydrogen is 13.6 ev and has maximal cross-section σ = 0.6 × 10−16 cm2 at 50 ev. full ionization energy for helium is 54.4 ev and its maximal cross-section σ = 0.3 × 10−16 cm2 is near 100 ev [17]. 5.3. ions yield estimation for the laser pulses with duration 20 ps the emitted charge 100 pc could be expected. with optimal electron temperature and for h2 concentration in range from 109 cm−3 to 1013 cm−3 [18] in a volume similar to current used pig ion source 2 cm3, from current density 18 a/cm2 there can expected from 102 to 105 positive ions. in normal operation the ratio between protons and h– ions is 10−2. very low probability of formation h– ion in the single pulse can be expected. even if the production ratio is lowered by two orders to 10−4, still up to 10 h– ions per volume could be expected. with considering a fact, than an exact h– formation process understanding is still in progress, an influence of more factors or processes, such as a wall biasing, could play its role. from this reason a formation of an h– ion in the single discharge bunch should not be rejected in advance. 5.4. discharge conditions and ions extraction after formation of the electron bunch by an irradiation of the photo-cathode, a multiple side effects can be expected. one from these effects could be the above mentioned parasitic photo-emission. other important effect will be positive ion bombardment of the cathode. parameter which can highly affect performance of the device will be a lifetime of produced ions which will probably reduce a minimal length of the ion pulse. 6. beam time distribution the low current regime of h– ions with energy 36.9 mev has its cyclotron frequency 25.8 mhz. this main carrier frequency is modulated by a 150 hz signal. particles in the cyclotron are accelerated with each half period and are caught in phase angle of about from −30° to 30° of this half period. the ions are revolving in approx. 6 ns micro-bunches with 32 ns spaces between them. with 10 % carrier modulation these micro-bunches are formed into 0.67 ms long bunches with 6 ms spaces. a simulation of the cyclotron response to a 110 ps short input pulse with 330 ions was made in the code durycnm4 [7, 19]. the ions from one input pulse are accelerated for 503–507 periods and divided into five output micro-bunches. output signal than lasts for 160 ns. energy range of the particles is from 36.75 to 36.9 mev. this cyclotron impulse response is made by a phase instabilities and space distribution of the micro-bunch, which affects energy increment difference particle gains each turn and leads to energy spread of the beam. because the extraction is made on the desired energy, each particle from the input bunch needs a different number of the accelerating turns. as a result on the output, the signal is time dispersed. this dispersion is coming from the machine nature and cannot be avoided by shortening of the input pulse. number of ions in individual micro-bunches are shown in fig. 10. 54 vol. 4/2016 ion source for a single particle accelerator figure 10. simulated u-120m impulse response. 7. conclusions this paper proposes a method which possibly could be used to produce a very small amount of the ions. considering many variable parameters that can be used to tune the ion source, it can be expected that the ions could be produced in countable amounts and delivered to the accelerating process on the request. the benefits of the presented method is mainly in the controlled production of the selected charge state and the number of the generated ions. as the demand of low current ion beams is increasing rapidly, overcoming the difficulties with generation the ions by the laser induced photo-emission could be worth the effort. references [1] g. a. drexler, c. siebenwirth, s. e. drexler, et al. live cell imaging at the munich ion microbeam snake – a status report. radiation oncology 10(1):42, 2015. doi:10.1186/s13014-015-0350-7. 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[6] t. vanát, j. pospíšil, f. křížek, et al. a system for radiation testing and physical fault injection into the fpgas and other electronics. in 2015 euromicro conference on digital system design (dsd), pp. 205–210. 2015. doi:10.1109/dsd.2015.98. [7] nuclear physics institute of the cas. mathematical simulation and analysis of the cyclotron u-120m, 2010. http://accs.ujf.cas.cz/index.php/en/component/ content/article/8. [8] the national institute of standards and technology. nist ebit introduction. http://www.nist.gov/pml/div684/grp01/intro.cfm. [9] lawrence livermore national laboratory. electron beam ion trap (ebit). llnl-web-409815 rev. 1, https://www-pls.llnl.gov/?url=about_ pls-physics_division-hed_physics-ebit. [10] highly charged ion trapping facility, gsi. the sparc ebit. http://web-docs.gsi.de/~stoe_exp/ laboratory/environment/ebit/ebit.php. [11] b. dunham. introduction to electron guns for accelerators. cornell university lecture notes, http://www.lepp.cornell.edu/~hoff/lectures/08s_ 688/08s_688_080225.pdf. [12] j. teichert, a. arnold, j. stephan, et al. progress of the rossendorf srf gun project. in proceedings of the 27th international free electron laser conference, pp. 534–537. 2005. [13] f. le pimpec, c. j. milne, c. p. hauri, f. ardana-lamas. quantum efficiency of technical metal photocathodes under laser irradiation of various wavelengths. applied physics a 112(3):647–661, 2013. doi:10.1007/s00339-013-7600-z. [14] c. w. schmidt. review of negative hydrogen ion sources. in proceedings of the linear accelerator conference, albuquerque, new mexico, usa, pp. 259–263. 1990. [15] m. p. stockli. volume and surface-enhanced volume negative ion sources. in proceedings of cas-cern accelerator school: ion sources, pp. 265–284. 2013. doi:10.5170/cern-2013-007.265. [16] d. p. moehs, j. peters, j. sherman. negative hydrogen ion sources for accelerators. ieee transactions on plasma science 33(6):1786–1798, 2005. doi:10.1109/tps.2005.860067. [17] physical measurement laboratory, nist. electron-impact ionization cross sections. http://physics.nist.gov/cgi-bin/ionization/ion_ data.php?id=hi&ision=i&initial=&total=y. [18] m. bacal, a. hatayama, j. peters. volume production negative hydrogen ion sources. ieee transactions on plasma science 33(6):1845–1871, 2005. doi:10.1109/tps.2005.860069. [19] m. čihák, o. lebeda, j. štursa. beam dynamic simulation in the isochronous cyclotron u-120m. in cyclotrons and their applications, eighteenth international conference, pp. 385–387. 2007. http://accelconf.web.cern.ch/accelconf/c07/ papers/385.pdf. 55 http://dx.doi.org/10.1186/s13014-015-0350-7 http://dx.doi.org/10.1088/0031-9155/56/8/006 http://dx.doi.org/10.1088/1674-1137/38/8/087003 http://dx.doi.org/10.1118/1.3469562 http://dx.doi.org/10.1088/0954-3899/41/8/087002 http://dx.doi.org/10.1109/dsd.2015.98 http://accs.ujf.cas.cz/index.php/en/component/content/article/8 http://accs.ujf.cas.cz/index.php/en/component/content/article/8 http://www.nist.gov/pml/div684/grp01/intro.cfm https://www-pls.llnl.gov/?url=about_pls-physics_division-hed_physics-ebit https://www-pls.llnl.gov/?url=about_pls-physics_division-hed_physics-ebit http://web-docs.gsi.de/~stoe_exp/laboratory/environment/ebit/ebit.php http://web-docs.gsi.de/~stoe_exp/laboratory/environment/ebit/ebit.php http://www.lepp.cornell.edu/~hoff/lectures/08s_688/08s_688_080225.pdf http://www.lepp.cornell.edu/~hoff/lectures/08s_688/08s_688_080225.pdf http://dx.doi.org/10.1007/s00339-013-7600-z http://dx.doi.org/10.5170/cern-2013-007.265 http://dx.doi.org/10.1109/tps.2005.860067 http://physics.nist.gov/cgi-bin/ionization/ion_data.php?id=hi&ision=i&initial=&total=y http://physics.nist.gov/cgi-bin/ionization/ion_data.php?id=hi&ision=i&initial=&total=y http://dx.doi.org/10.1109/tps.2005.860069 http://accelconf.web.cern.ch/accelconf/c07/papers/385.pdf http://accelconf.web.cern.ch/accelconf/c07/papers/385.pdf acta polytechnica ctu proceedings 4:50–55, 2016 1 introduction 2 beam current limitation 2.1 u-120m pig ion source 2.2 pulsed pig ion source 3 inspiring ion sources 3.1 ebis/ebit ion source 3.2 rf gun 4 laser induced photo-emission ion source 4.1 photo-cathode 5 ions production 5.1 negative hydrogen ion formation 5.2 positive ions formation 5.3 ions yield estimation 5.4 discharge conditions and ions extraction 6 beam time distribution 7 conclusions references 34 acta polytechnica ctu proceedings 1(1): 34–37, 2014 34 doi: 10.14311/app.2014.01.0034 a novel approach in the wimp quest: cross-correlation of gamma-ray anisotropies and cosmic shear marco regis1 1dipartimento di fisica, università di torino and infn, torino, italy corresponding author: regis@to.infn.it abstract we present the cross-correlation angular power spectrum of cosmic shear and gamma-rays produced by the annihilation/decay of weakly interacting massive particle (wimp) dark matter (dm), and by astrophysical sources. we show that this observable can provide novel information on the composition of the extra-galactic gamma-ray background (egb), since the amplitude and shape of the cross-correlation signal depend on which class of sources is responsible for the gamma-ray emission. if the dm contribution to the egb is significant (at least in a definite energy range), although compatible with current observational bounds, its strong correlation with the cosmic shear (since both signals peak at large halo masses) makes such signature potentially detectable by combining fermi-lat data with forthcoming galaxy surveys, like dark energy survey and euclid. keywords: dark matter extragalactic gamma rays background weak gravitational lensing. 1 introduction the presence of gravitational anomalies at different scales (galactic, cluster, large scale structure, and cosmological scales) is observationally well-established. however, it is not fully understood yet whether such anomalies are due to a new form of matter (i.e., dark matter (dm)) or to a new form of interaction (e.g., a modification of the laws of gravity). currently, the dm solution represents the mainstream, mainly because there is no theory of gravity which can account for all the anomalies at different scales simultaneously. for example, mond (modified newtonian dynamics) is a theory stating that acceleration is not linearly proportional to gravitational force at small accelerations, and seems to be quite successful in explaining galactic rotation curves. however, it cannot explain cluster dynamics without the addition of some dm components. on the other hand, it is possible that we just haven’t been good enough to build the proper extension to newtonian and einstein gravity. moreover, if dm interacts only gravitationally with ordinary matter, the two classes of solution are hardly distinguishable. one of the best ways to settle the issue would be to detect a non-gravitational signal of dm coming from the regions with gravitational anomalies. in this talk, we will present a novel strategy in this direction. in particular, if dm is made by weakly interacting massive particles (wimps), they have weak but non-negligible interactions with ordinary matter, and one of the predictions of the model is sizable fluxes of γ-rays from dm halos. both cosmic shear and cosmological gamma-ray emission stem from the presence of dm in the universe: dm structures are responsible for the bending of light in the weak lensing regime and while γ-rays can be produced by astrophysical sources hosted by dm halos (i.e. star-forming galaxies (sfg) or active galactic nuclei (agn)), dm itself may be a source of γ-rays, through its self annihilation or decay, depending on the properties of the dm particle. those γ-rays emitted by dm should therefore have the potential to exhibit strong correlation with the gravitational lensing signal. the most recent measurement of the egb was performed by the fermi-lat telescope in abdo et al. (2010)a, covering a range between 200 mev and 100 gev: the emission is obtained by subtracting the contribution of resolved sources (both point-like and extended) and the galactic foreground (due to cosmic rays interaction with the interstellar medium) from the whole fermi-lat data. unresolved astrophysical sources like blazars or radio galaxies contribute to the egb but the exact amount of their contribution is still unknown. the γ-rays produced by dm annihilation or decay can also contribute to egb (e.g., ullio et al. (2002)). however, the fact that the egb energy spectrum is compatible with a power-law, without any evident spectral feature, suggests that dm cannot play a leading role in the whole energy range, see abdo et al. (2010)b. 34 http://dx.doi.org/10.14311/app.2014.01.0034 a novel approach in the wimp quest: cross-correlation of gamma-ray anisotropies and cosmic shear in the angular anisotropies of the egb emission, the dm also plays a subdominant role: indeed, a detection of a significant auto-correlation angular ps has been recently reported in ackermann et al. (2012) (for multipoles ` > 100, which is the range of interest for our analysis, since there the contamination of the galactic foreground can be neglected), but the features of such a signal (in particular its independence on multipole and energy) seem to indicate an interpretation in terms of blazars, harding & abazajian (2012). both cosmic shear and γ-ray emission depend on the large scale structure of the universe: because this is what generates the lensing effect and because those same structures can produce γ-rays. here we will show how to use the lensing signal in order to possibly disentangle a dm signal in the egb. a detection of such cross-correlation would demonstrate that the weak lensing observables are indeed due to particle dm matter and not to possible modifications of general relativity. this presentation is based on camera et al. (2012). figure 1: left: egb emission as a function of observed energy for the four extragalactic components described in the text. data are from abdo et al. (2010)a. right: γ-ray angular ps at e > 1 gev for the same models of the left panel. the observed angular ps is summarized by the black band from ackermann et al. (2012). this figure is taken from camera et al. (2012). figure 2: left: window functions vs. redshift. for γ-ray sources we consider the flux above 1 gev normalized to the total egb intensity measured by fermilat. right: three-dimensional ps of cross-correlation shear/γ-rays at z = 0. this figure is taken from camera et al. (2012). 2 theoretical modeling the source intensity along a given direction ~n can be written as: ig(~n) = ∫ dχg(χ,~n) w̃(χ) , (1) where χ(z) is the radial comoving distance, g is the density field of the source, and w̃ is the window function (which does not depend on ~n). we then define a normalized version w = 〈g〉w̃ , so 〈ig〉 = ∫ dχ w(χ). expanding the intensity fluctuations of two source populations i and j in spherical harmonics, one can compute the cross-correlation angular ps (here in the dimensionless form): c (ij) ` = 1 〈ii〉〈ij〉 ∫ dχ χ2 wi(χ) wj(χ)pij(k = `/χ,χ) . (2) the definition of the 3-dimensional ps pij is 〈f̂gi (χ,~k)f̂ ∗ gj (χ′,~k′)〉 = (2π)3δ3(~k −~k′)pij(k,χ,χ′) , (3) where fg ≡ [g(~x|m,z)/ḡ(z) − 1] (f̂g is its fourier transform) and the limber approximation (k = `/χ) is assumed to hold. we consider the sources to be characterized by a parameter m (typically the mass), and g(~x|m) is the density field of an object associated to m, while ḡ(z) = 〈g(~n,z)〉. pij can be computed following the so-called halo-model approach. the two-point correlation is given by the sum of two components, the 1-halo and 2-halo terms, i.e. pij = p 1h ij +p 2h ij , (scherrer & bertschinger (1991), ando & komatsu (2006)): p1hij (k) = ∫ dm dn dm f̂∗i (k|m) f̂j(k|m) (4) p2hij (k) = [∫ dm1 dn dm1 bi(m1)f̂ ∗ i (k|m1) ] × [∫ dm2 dn dm2 bj(m2)f̂j(k|m2) ] p lin(k) ,(5) where dn/dm is the number density distribution of sources, p lin is the linear matter ps, and bi(m) is the linear bias between the object i and matter. note that the average of 〈g〉 is given by: ḡ(z) = 〈g(~n,z)〉 = ∫ dm dn dm ∫ d3~xg(~x|m,z) , (6) which implies that at small k (where f̂ ∼∫ d3~xg(~x|m)/ḡ) the terms in the square-brackets in eq. (5) are ∼ 1 (except in the case of a significant bias). 35 marco regis the 2-halo term is thus normalized to the standard linear matter ps at small k, which motivates the normalization of the window function introduced above. we aim at cross-correlating the shear signal (source i in eqs. (2–5)) with γ-rays emitted by dm, sfgs, and blazars (source j in eqs. (2–5)). 3 results the above formalism can be applied to all the mentioned components, by considering the specific g, w , and dn/dm of each case. for the sake of clearness we will focus on a benchmark model for each component. they are described in camera et al. (2012). in particular, we note that, although we take dm to provide a significant contribution of the egb at e ≥ 1 gev in fig. 1a, it is basically impossible to obtain an evidence for dm from the angular ps of γ-rays alone because the latter is dominated by the blazar contribution. the contributions to the egb and to the γ-ray autocorrelation aps of each components are shown in fig. 1. in fig. 2 we show the ingredients of eq. (2) for the computation of the shear/γ-ray cross-correlation angular ps: the window functions and 3d power spectra. the window function for the cosmic shear signal nicely overlaps with the dm window function, both for annihilating and decaying dm, while this happens only at intermediate redshifts for the sfg window function and only at high redshifts for the case of blazars. this suggests that a tomographic approach could be a powerful strategy to further disentangle different contributions in the angular ps (this will be explored in a forthcoming work, camera et al. (in preparation)). the shear signal is stronger for larger dm masses. the same is true also for the γ-ray signal from dm and this fact gives a large 1-halo contribution which dominates in the range k = 1−10 h/mpc in fig. 2b. those wavenumbers correspond to ` ' 100 − 1000. galaxies have masses ≤ 1014m�, thus they correlate with the shear signal of lower-mass halos and the 1-halo contribution becomes important at smaller scale k � 1h/mpc. thus, an important result is that, since both the shear and dm-induced γ-ray signals are stronger for larger halos, their cross-correlation is more effective with respect to the case of astrophysical sources at intermediate wavenumbers. this, together with the sizable overlapping of the dm γ-ray and shear window functions at low redshift, leads to the expectation of a sizable dm signal in the angular ps, which is indeed what we find in fig. 3. the observational forecasts for the cross-correlation between dark energy survey (des), abbott et al. (2005), or euclid, laureijs et al. (2011), and fermilat are shown for the benchmark models considered in this work (for error estimates, we take observational performances from atwood et al. (2009), laureijs et al. (2011), and abbott et al. (2005). fig. 3 shows that a dm signal can be disentangled in the angular ps at ` ≤ 103. the same conclusion can be derived for dm models with different mass and annihilation/decay channels, provided the dm is a significant component of the total γ-ray egb (at least in one energy bin) as in our assumptions. figure 3: left: cross-correlation between cosmic shear and γ-ray emission, for the different classes of γray emitters described in the text (with a γ-ray threshold expected for fermi-lat after 5 years of exposure). each contribution is normalized by multiplying eq. (2) by 〈ij〉/〈iegb〉 to make them additive. des is taken as the reference galaxy survey. error bars are estimated for the total signal (in black). right: same as in the left panel but for annihilating dm, with euclid as the reference galaxy survey. this figure is taken from camera et al. (2012). 4 conclusions in this talk, we discussed the cross-correlation angular power spectrum of weak lensing cosmic shear and γrays produced by wimp annihilations/decays and astrophysical sources. we showed that this method can provide novel information on the composition of the egb. since the shear signal is stronger for structures of larger masses and most of the γ-ray emission from decaying and annihilating dm is also produced in large mass halos, their cross-correlation is typically stronger than the case of astrophysical sources (which are associated to galactic-mass halos) for ` ' 100 − 1000. the combination of fermi-lat with forthcoming surveys like des and euclid can thus potentially provide evidence for wimps. acknowledgement we would like to thank the organizers for the invitation to the workshop and the participants for interesting discussions. we acknowledge infn grant fa51. 36 a novel approach in the wimp quest: cross-correlation of gamma-ray anisotropies and cosmic shear references [1] t. abbott et al. [des collab.], astro-ph/0510346. [2] a. abdo et al. [fermi-lat collab.], phys. rev. lett. 104 (2010) 101101. doi:10.1103/physrevlett.104.101101 [3] a. abdo et al. [fermi-lat collab.], jcap 1004 (2010) 014. [4] m. ackermann et al. [fermi lat collab.], phys. rev. d 85 (2012) 083007. doi:10.1103/physrevd.85.083007 [5] s.’i. ando, e. komatsu, phys. rev. d73 (2006) 023521. [6] w. atwood et al. [lat collab.], ap. j. 697 (2009) 1071. doi:10.1088/0004-637x/697/2/1071 [7] s. camera, m. fornasa, n. fornengo and m. regis, astrophys. j. 771 (2013) l5 [arxiv:1212.5018 [astro-ph.co]]. [8] s. camera, m. fornasa, n. fornengo and m. regis, in preparation. [9] j. harding and k abazajian, jcap 1211 (2012) 026. [10] r. laureijs et al., arxiv:1110.3193; l. amendola et al. [euclid theory working group collab.], arxiv:1206.1225; http://www.euclid-ec.org. [11] r. scherrer and e. bertschinger, ap. j. 381, 349 (1991). [12] p. ullio et al. phys. rev. d 66 (2002) 123502. doi:10.1103/physrevd.66.123502 discussion sergio colafrancesco: what are the effects of dm-density profiles and of the cross-correlation of dm halos with gamma-ray emitting galaxies on the results you have shown? marco regis: dm density profiles in the inner part are model dependent; however this affects the cross-correlation angular power spectrum only at multipoles ` > 103−104, so above the range of our interest. since the bulk of anisotropies from dm comes from very large halos, the galactic mass halos are not crucial in the estimate of the total dm signal. jim beall: how much of an improvement in the error bars do you need to better delimit the model? marco regis: this changes with the overall normalization of the wimp cross-correlation signal which depends on the (unknown) interaction rate. however, the bottom-line is that, provided the dm is one of the most relevant components of the egb in at least a energy band, current and next future experiments, like fermi-lat and des, would suffice. 37 http://dx.doi.org/10.1103/physrevlett.104.101101 http://dx.doi.org/10.1103/physrevd.85.083007 http://dx.doi.org/10.1088/0004-637x/697/2/1071 http://dx.doi.org/10.1103/physrevd.66.123502 introduction theoretical modeling results conclusions 265 acta polytechnica ctu proceedings 1(1): 265–268, 2014 265 doi: 10.14311/app.2014.01.0265 first hints of pressure waves in a helical extragalactic jet: s5 0836+710 manel perucho1 1departament d’astronomia i astrof́ısica. universitat de valència. c/dr. moliner 50, 46100, burjassot, valencian country, spain. corresponding author: manel.perucho@uv.es abstract one of the open questions in extragalactic jet astrophysics is related to the nature of the observed radio jet, namely whether it traces a pattern or the flow structure itself. in this paper i summarize the evidence collected for the presence of waves in extragalactic jets. the evidence points towards the peak of emission in helical jets corresponding to pressuremaxima of a wave that is generated within the core region and propagates downstream. making use of a number of very long baseline interferometry (vlbi) observations of the radio jet in the quasar s5 0836+710 at different frequencies and epochs, perucho et al. (2012a) were able to observe wave-like behavior within the observed radio-jet. the ridge-line of the emission in the jet coincides within the errors at all frequencies. moreover, small differences between epochs at 15 ghz reveal wave-like motion of the ridge-line transversal to the jet propagation axis. the authors conclude that the helicity is a real, physical structure. i report here on those results and discuss them in the light of new results recently announced by other authors that confirm the presence of waves in the close-by object bl lac (cohen et al., in preparation). keywords: galaxies: jets hydrodynamics instabilities quasars: individual: s5 0836+710. 1 introduction the parsec-scale structure of jets in active galactic nuclei (agn) is mainly observed in the radio band, using the vlbi technique. the nature and properties of the emitting region as related to the flow are still scarcely known. jets can be interpreted as flows because the larmor radius of particles is very small when compared to the spatial scales of the problem studied (blandford & rees 1974). this interpretation has different implications: such systems, which are composed of magnetic fields and particles propagating along the jet channel are expected to host the growth of different hydrodynamical and/or magnetic instabilities. the instabilities have been claimed to be the cause of many of the structures observed (knots, bendings, helices; see e.g., reviews in hardee 2006, 2011, perucho 2012). the instabilities are waves triggered by any external or internal perturbation that grow in amplitude with distance, as they are advected with the flow. thus, if instabilities do really grow in jets, we would expect to observe wave-like structures and motions in extragalactic jets. perucho et al. (2012a) used observations of the jet in s5 0836+710 at different frequencies and epochs and showed that theridge line of this jet behaves as expected if it is interpreted as a pressure wave. aaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaa aaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaa aaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaa the luminous quasar s5 0836+710 is at a redshift z = 2.16 and hosts a powerful radio jet extending up to kiloparsec scales (hummel et al. 1992). at this redshift, 1 mas ' 8.4 pc (see, e.g., mojave database). vlbi monitoring of the source showed kink structures (krichbaum et al. 1990) and yielded estimates of the bulk lorentz factor γj = 12 and the viewing angle αj = 3 ◦ of the flow at milliarcsecond scales (otterbein et al. 1998). the jet was observed at 1.6 and 5 ghz with vsop (vlbi space observatory program, a japanese-led space vlbi mission), and oscillations of the ridge-line were also observed (lobanov et al. 1998, 2006). these oscillations were interpreted as kelvin-helmholtz (kh) instabilities (lobanov et al. 2006, perucho & lobanov 2006). the source has also been monitored by the mojave program (lister et al. 2009). finally, it has been suggested that the lack of a collimated jet structure and hot-spot at the arcsecond scales is due to jet disruption, as indicated by the growth in the amplitude of the helical structure (perucho et al. 2012b). 265 http://dx.doi.org/10.14311/app.2014.01.0265 manel perucho 2 evidence for waves in the jet of s5 0836+710 here i summarize the evidence that points to the presence and development of waves in the jet of s5 0836+710, as given by perucho et al. (2012a). the different observations of the jet at different frequencies ranging from 1.6 ghz to 43 ghz, for a given epoch, give the same ridge-line positions within the errors. it needs to be determined whether the small scale ripples in the ridge-line are caused by uv-coverage effects at the different epochs/frequencies, which is plausible. however, the large scale motion coincides for all frequencies and this can only happen if the observed structure corresponds to a physical one. in addition, high resolution images at 15 ghz confirm that the ridgr-line position does not necessarily coincide with the centre of emission of the radio jet, as expected from an asymmetric pressure maximum triggering a helical instability in the jet. within the first milliarcsecond, high resolution images of the jet allowed the measurement of transversal displacements, which show a clear wave-like oscillation pattern with distance. the obtained ridge-line velocities at these scales are superluminal, but this is due, according to the authors, to: 1) relativistic wave velocities, 2) a small-scale oscillation of the core, very difficult to detect in this distant source, but observed in others (e.g., m 81, see mart́ı-vidal et al. 2011), and 3) displacements of the ridge-line being smaller than the errors in the determination of the position. the amplitude of the oscillation grows along the jet propagation direction, which can be due to the coupling to an unstable mode of the kh or current-driven (cd) instabilities. the evidence collected in the previous paragraph led the authors to conclude that the ridge-line of the jet is tracing a pressure-maximum of a wave within the jet, and that even the radio-core is subject to the oscillation produced by the wave. recently, cohen et al. (2013) presented new results of the nearby source bl lac (in the conference the innermost regions of relativistic jets and their magnetic fields, held in granada 10th-14th june, 2013), using long-term monitoring of the source at 15 ghz (cohen et al. in preparation), which confirm that the ridge-line of emission within the jet shows a wave-like behavior. in this case, the authors claim that the oscillation of the ridge-line is caused by a strong magnetic field. in addition, grossberger et al. (2012, and private communication) have followed the ejection and evolution of components in the broad-band radiogalaxy 3c 111 and shown that the position angles of the injected components vary with time and fill a wide channel after several years. this channel is wider than the observed jet at each epoch. this result also points to an oscillation of the bright regions of the jet (necessarily the same as that identified with the ridge-line in other works) within a wide channel in which the observed radio jet is embedded. 3 discussion 3.1 on the nature of the waves given the evidence found for the presence of waves in jets, we can now ask which kind of waves these are and what is their origin. regarding the latter, it is still unclear whether the waves show any regular periodical pattern or not, but the long-term oscillations could be attributed to periodical or quasi-periodical processes in the injection region, such as precession. shorter-term oscillations can be attributed to perturbations triggered by asymmetries in the jet itself or in the ambient medium, lateral winds or entrainment of clouds that rotate around the galactic centre from a side of the jet, for instance. however, all these processes are extremely difficult to detect and isolate. as far as the nature of the waves, there is (already) open debate as to whether they are magnetic or purely hydrodynamical waves. the claim that the waves correspond to magnetic instabilities is based on the fact that the observed radio components seem to follow the ridgeline as they propagate (cohen et al., in preparation): if the jet ridge line is identified with a kink instability in the magnetic field of a magnetically dominated jet, the flow would be forced to follow this kink. however, let us consider the possibility that the ridge-line corresponds to the pressure maximum produced by the kink (e.g., mizuno et al. 2011). this implies that within a given jet cross-section we can find a high-pressure region and a low-pressure one. in this case, the passage of a ballistic component following a straight path along the jet, could enhance the emission of the whole jet, but the effect would be more visible on the ridge-line (high-pressure region). as a result, any radio-component fitted to the region would be centered there. perucho et al. (2006) showed, via numerical simulations, that an asymmetric perturbation at the base of a relativistic jet propagates downstream and naturally creates a pressure maximum helix, which can eventually distort the jet flow and force it into a helical path. in this view, injected components associated with perturbations in flow density at injection that produce shock waves, end up following the helical path when they have dissipated most of their kinetic energy. previous to this, they propagate ballistically. a long-standing cross-debate about jet physics is related to the ballistic nature of the flow in apparently helical radio jets. this point again introduces a difference in behavior, but still not a way to distinguish, 266 first hints of pressure waves in a helical extragalactic jet: s5 0836+710 between a kinetically dominated flow and a magnetically dominated one. if the jet is strongly magnetized, the flow will always follow the field lines and will thus have a twisted path if the field is helical, whereas in the kinetically dominated case, the flow could follow a straight stream-line even if the radio-jet shows a helical structure (because it corresponds to the bright high pressure region within a wide channel). this is true while the amplitude of the instability is not large, but when it grows to nonlinear values, the jet flow starts to deviate from its original trajectory by the pressure gradients within the jet (see, e.g., nakamura & meier 2004, mizuno et al. 2011, perucho et al. 2005, 2006 for the different cases). thus, a purely hydrodynamical jet could produce observational patterns similar to a strongly magnetized jet, and further work is needed to understand the physics behind the ridge-line dynamics. moreover, even if no perturbations are injected in the jet, i.e., the flow is close to stationary, then the fitted components in the radio jet would all be associated with the ridge-line (see, e.g., hardee 2003), just because it corresponds to the brightest regions within the jet. if the ridge-line corresponds to a propagating wave, then the fitted-components would also move along the jet and, of course, on top of the ridge-line. it is thus crucial to disentangle whether the components correspond to waves propagating along the jet and whether or not they follow the ridge-line for understanding its nature. 3.2 the case of s5 0836+710 the wave associated with the ridge-line could couple to a growing instability, as indicated by the growth in amplitude of the helix with distance in the case of s5 0836+710. it is difficult to distinguish between current-driven (cd) instability and kh instability modes, both being solutions to the linearized relativistic, magnetized flow equations and both being possible sources of helical patterns. recent work on cd instability by mizuno et al. (2011) shows that a helical kink propagates with the jet flow if the velocity shear surface is outside the characteristic radius of the magnetic field, i.e., approximately the radius at which the toroidal magnetic field is a maximum. if the observed pattern corresponds to a cd kink instability, the observed transversal oscillation in the jet of 0836+710 requires that the kink be moving with the flow and implies that the transversal velocity profile is broader than the magnetic field profile, i.e., the velocity shear surface lies outside the characteristic radius of the magnetic field. thus, in this case the jet would have a magnetized spine surrounded by a particle dominated outer region and a cd kink in the spine would move with the flow.. the measurement of the opening angle reveals a very similar value at all frequencies for which we have significant measurements, which would favor the radio-jet tracing the whole channel, because the coincidence in the opening angles is less probable in the case of the radio jet corresponding to different regions across the jet. thus, it seems that the radio emission is generated in the same region across the jet at all frequencies and that the jet flow indeed follows a helical path. as stated above, non-ballistic motion is compatible with a magnetically dominated flow, but also with the development of kh-instability modes to large amplitude, because they can dissipate kinetic energy into internal energy and decelerate the flow. it has been reported that the development of the helical kh instability can force the jet flow into a helical path, albeit different from the helicity of the pressure maximum, the latter showing larger amplitude (hardee 2000). the larger the lorentz factor of the flow and the shorter the wavelength of the mode, the more different is the helicity of the flow as compared to that of the pressure maximum. this can be understood in terms of the larger inertia of the jet flow with increasing lorentz factor (hardee 2000, perucho et al. 2005). in summary, the fact that the flow is non-ballistic when the amplitude of the perturbation is large does not necessarily imply that the flow is magnetically dominated and that the waves are magneto-sonic waves. 3.3 the core one other important consequence of the work presented in perucho et al. (2012a) is the possible small-scale oscillation of the radio-core around an equilibrium position. this effect has been reported by mart́ı-vidal et al. (2011) for m 81, using the phase-referencing technique, and it could have important implications for the nature of the radio-core. 4 conclusions it is possible to associate the ridge-line of emission of helical jets to wave patterns tracing a (thermal or magnetic) pressure maximum produced by the growth of instabilities. the core of the jet could well also be oscillating and this motion should be tested and corrected in order to study the parsec-scale kinematics of the radio-components in jets. further combined observational and theoretical studies like this one are required to get more information on the nature of the growing instability and on the properties of the jet flow. acknowledgement i acknowledge financial support by the spanish “ministerio de ciencia e innovación” (micinn) grants aya2010-21322-c03-01 and aya2010-21097-c03-01. i 267 manel perucho also acknowledge p.e. hardee, j.m. mart́ı, m. cohen, and d.l. meier for interesting discussions on the topic. references [1] blandford, r.d., rees, m.j.: 1974, mnras 169, 395. doi:10.1093/mnras/169.3.395 [2] cohen, m.h., et al.: 2013, in the innermost regions of relativistic jets and their magnetic fields. http://jets2013.iaa.es/sites/jets2013.iaa. es/files/imagecache/granada may2013 v3.ppt [3] grossberger, c., kadler, m., wilms, j., et al.: 2012, acpol 52, 18. [4] hardee, p.e.: 2000, apj 533, 176. doi:10.1086/308656 [5] hardee, p.e.: 2003, apj 597, 798. doi:10.1086/381223 [6] hardee, p.e.: 2006, in proceedings of relativistic jets: the common physics of agn, microquasars and gamma-ray bursts, eds.: p.a. hughes and j.n. bregman, aip conference proceedings 856, 57. [7] hardee, p.e.: 2011, in proceedings of iau symposium 275: jets at all scale, eds.: g. romero, r. sunyaev and t. belloni, iau conference series 275, 41. [8] hummel, c.a., muxlow, t.w.b., krichbaum, t.p., et al.: 1992, a&a 266, 93. [9] krichbaum, t.p., hummel, c.a., quirrenbach, a., et al.: 1990, a&a 230, 271. [10] lister, m., aller, h.d., aller. m.f., et al.: 2009, aj 137, 3718 [11] lobanov, a.p., krichbaum, t.p., witzel, a., et al.: 1998, a&a 340, 60. [12] lobanov, a.p., krichbaum, t.p., witzel, a., zensus, j.a.: 2006, pasj 58, 253. [13] mart́ı-vidal, i., marcaide, j.m., alberdi, a.: 2011, a&a 533, a111. [14] mizuno, y., hardee, p.e., nishikawa, k.-i.: 2011, apj 734, 19. doi:10.1088/0004-637x/734/1/19 [15] nakamura, m., meier, d.l.: 2004, apj 617, 123. doi:10.1086/425337 [16] otterbein, k., krichbaum, t.p., kraus, a., et al.: 1998, a&a 334, 489. [17] perucho, m.. mart́ı, j.m., hanasz, m.: 2005, a&a 443, 863. [18] perucho, m., lobanov, a.p., mart́ı, j.m., hardee, p.e.: 2006, a&a 456, 493. [19] perucho, m., lobanov, a.p.: 2007, a&a 469, l23. [20] perucho, m.: 2012, in proceedings of high energy phenomena in relativistic outflows iii (heproiii), eds. j.m. paredes, m. ribó, f.a. aharonian and g.e. romero, ijmps 8, 241. [21] perucho, m., kovalev, y.y., lobanov, a.p., hardee, p.e., agudo, i.: 2012a, apj 749, 55. doi:10.1088/0004-637x/749/1/55 [22] perucho, m., mart́ı-vidal, i., lobanov, a.p., hardee, p.e.: 2012b, a&a 545, a65. discussion sergio colafrancesco: is there any radio polarization measurement that can be used to better describe the helical structure of the jet and, if so, what are the constraints that can be set on the acceleration mechanisms at the base of the jet? manel perucho: krichbaum et al. (1990) reported polarization measurements and claim that the field is aligned with the jet axis. if their interpretation is correct, it would be difficult to attach the helical morphology of the jet to the magnetic field. nevertheless, new polarization measurements should be performed to check whether the magnetic field plays any crucial role in the generation of the helical structure. regarding your second question, it is difficult to observe the acceleration region in this jet, but we have evidences of component acceleration in other agn jets, included in the mojave sample. herman marshall: if the perturbations are unstable and disrupt the jet at kpc scales, how do you explain that most jets seem to propagate to tens or hundreds of kpc? manel perucho: in that sense, this is a unique object, because it is classified as an frii in terms of its radio luminosity, but its morphology denies (see perucho et al. 2012b). this is one of the reasons why we think that, in this sole case, the jet could be destroyed by the growth of the instability, unlike all other known frii jets for which we can observe the kiloparsec scale structure. 268 http://dx.doi.org/10.1093/mnras/169.3.395 http://dx.doi.org/10.1086/308656 http://dx.doi.org/10.1086/381223 http://dx.doi.org/10.1088/0004-637x/734/1/19 http://dx.doi.org/10.1086/425337 http://dx.doi.org/10.1088/0004-637x/749/1/55 introduction evidence for waves in the jet of s5 0836+710 discussion on the nature of the waves the case of s5 0836+710 the core conclusions 60 acta polytechnica ctu proceedings 2(1): 60–65, 2015 60 doi: 10.14311/app.2015.02.0060 on the influence of magnetic field on accretion processes in cvs d. v. bisikalo1, a. g. zhilkin1 1institute of astronomy ras, moscow, russia corresponding author: bisikalo@inasan.ru abstract we consider the influence of such parameters as the value of the proper magnetic field ba and the spin-rotation velocity of the white dwarf on accretion processes in cvs. the results of 3d mhd simulations have shown that the accretion rate is a non-monotonic function of ba: with growing ba it raises in the intermediate polars and decreases in the polars. the maximal accretion rate occurs in the systems, transiting from the stage of intermediate polars to polars; it’s value reaches ∼60% of the initially set mass transfer rate. we have also shown that the acretion rate decreases with the growing spin-rotation velocity of the white dwarf. keywords: cataclysmic variables intermediate polars polars accretion mhd simulations. 1 introduction two main classes of cvs where magnetic field significantly influences accretion processes are the intermediate polars and polars [1]. these are semi-detached binary systems, consisting of a low-mass late type star (donor-star) and a white dwarf (accretor). in the polars, white dwarfs possess strong magnetic fields (ba ≈ 107–108 g on the surface). observations show that no accretion disks form in polars. instead, the material, issuing from the l1 point, forms a collimated stream, that flows onto one of the magnetic poles of the accretor along magnetic field lines[1, 2]. in the intermediate polars, white dwarfs possess relatively weak magnetic fields (ba ≈ 104–106 g on the surface). these systems occupy an intermediate stage between the polars and non-magnetic cvs. we should note that in these systems the spin-rotation periods of the primaries may be significantly shorter than the orbital periods [3]. in our previous works [4, 5] we, by means of 3d mhd simulations, studied the flow structure of the intermediate polars and polars, having various surface magnetic fields of the accretors ba, ranging from 10 5 to 108 g. this allowed us to investigate conditions of accretion disk formation and to find a criterion, separating two types of the flow, corresponding to the intermediate polars and polars. besides, in [5] we also investigated how the spin-rotation velocity of the accretor influences the mhd flow structure in the system.in this paper we summarize the results of our previous research and show how the accretion rate varies, depending on two parameters as the magnetic induction on the surface of the white dwarf and its spin-rotation velocity. 2 what happens with the flow structure if ba grows to investigate how the value of the accretor’s surface magnetic field influences mhd flow structure in a close binary system we have conducted 7 computational runs with various ba: 10 5 g (model 1); 5 × 105 g (model 2); 106 g (model 3); 5 × 106 g (model 4); 107 g (model 5); 5 × 107 g (model 6); and 108 g (model 7). we suppose that the accretor’s magnetic field is of dipole type and the axis of the field is inclined with respect to the accretor’s rotation axis by θ = 30◦. for the computations we used our method of investigations of mass transfer processes in semi-detached binary systems where the accretors possess strong magnetic fields [4, 5]. the method is based on the idea that plasma dynamics is determined by the relatively slow average motion of plasma against a background of which very rapid mhd waves propagate. the equations, describing the slow motion of plasma, are derived via a procedure of averaging over the rapidly propagating pulsations. strong external magnetic field in this approach plays a role of an effective liquid interacting with plasma. in the equation of motion the mean electromagnetic force is analogous to the friction force working in multi-component plasma. in our numerical model we take into account processes of the magnetic field diffusion caused by the currents’ dissipation in turbulent vortices and magnetic buoyancy. besides, when averaging the induction equation over the rapidly propagating mhd waves we introduce an additional source of magnetic field dissipation (wave dissipation). in order to see how namely the magnetic induction influences the solution we fix all the other param60 http://dx.doi.org/10.14311/app.2015.02.0060 on the influence of magnetic field on accretion processes in cvs eters. the computations have been conducted for a close binary system, having parameters of the ss cygni (see [6]). the donor-star (red dwarf) in this system has the mass md = 0.56m� and the effective temperature of 4000 k. the accretor (white dwarf) has the mass ma = 0.97m� and the temperature of 37000 k. the orbital period of the system is porb = 6.6 hours and its binary separation is a = 2.05r�. we have supposed that the accretion disk forms in the system for the first time so the accretor’s spin-rotation is synchronous, i.e. pspin = porb. the influence of the accretor’s spinrotation on a formed accretion disk is considered in the next section. the obtained solutions can be divided into two groups. the first group contains models 1, 2, and 3 with relatively weak magnetic fields where accretion disks form. these models correspond to the intermediate polars. in the models of the second group (models 4–7) magnetic fields are strong and there are no accretion disks formed. these models represent the polars. the typical flow structure in the intermediate polars is shown in fig. 1, where we plotted an iso-surface of density logarithm lgρ = −4.5 (in units of ρ(l1)) obtained in the model 1 (ba = 10 5 g). in figure we also show the magnetic flow lines originating on the accretor’s surface, accretor’s rotation axis (thin straight line) and magnetic axis (bold inclined line). we note that all the gas dynamic structures found in our previous studies [7, 8, 5], including the accretion disk and shock waves exist in the obtained solution. near the accretor the magnetosphere forms and material is accreted via funnel flows. with growing magnetic field the outer radius of the disk significantly decreases. the efficiency of the magnetic breaking and angular momentum transfer increases. the region of the magnetosphere becomes significantly larger. finally, when ba = 10 6 g (model 3, fig. 2) the accretion disk almost degenerates. the material makes only 1–2 cycles before falling onto the accretor. to describe this flow structure a term ”spiralodisk” is more appropriate, since the velocity distribution in this structure strongly differs from the keplerian. the outer radius of this ”spiralo-disk” is approximately 0.1a and the inner radius is determined by the size of the magnetosphere. a significant part of the ”spiralo-disk” is accretion funnel flows. we can state that this solution is an ultimate case of the intermediate polars. to determine a criterion, separating the intermediate polars and polars let us consider behavior of the gas stream, issuing from the l1 point into the roche lobe of the accretor. taking into account that the gas in the stream moves with supersonic velocity we can consider its behavior using the ballistic approach and omitting the effects of pressure and magnetic field [9, 10]. figure 1: 3d flow structure for the model 1 with ba = 10 5 g. iso-surface of density (lgρ = −4.5 in units of ρ(l1)) and magnetic field lines are shown. the accretor’s rotation axis is shown by the thin straight line, magnetic axis — by the bold inclined line. figure 2: the same as in fig. 1 for the model 3 with ba = 10 6 g. figure 3: the same as in fig. 1 for the model 5 with ba = 10 7 g. 61 d. v. bisikalo, a. g. zhilkin 9 ,1 0 y e a r a m m e & , g a b 9 , 1 0 y e a r m m e & , g a b ,% a m m & & , g a b 9 ,1 0 y e a r a m m e & orb spin p p figure 4: the dependence of the accretion rate ṁa on value of the magnetic field ba (left row) in absolute values (top panel) and in % from the mass transfer rate ṁ (bottom panel). the vertical lines limit the specific amplitudes of the accretion rate variations. in the upper-right panel we show the mass transfer rate as a function of the value of magnetic field. in the lower-right panel we show the mass transfer rate as a function of the accretor’s spin velocity in the model with ba = 10 5 g. analysis of the trajectories [11] shows that the stream comes very close to the accretor’s surface at a distance rmin. if rmin is larger than the radius of the magnetosphere rm the magnetic field does not strongly influence the motion of material. the stream rounds the star and intersects itself at a certain point. the further evolution of this flow results in the formation of an accretion disk in the system. if rmin is smaller than rm then the magnetic field starts to strongly influence the stream at a certain part of its trajectory. the action of electromagnetic forces in this region slows the stream down and causes loss of the angular momentum. as a result the stream can not round the star and form an accretion disk. thus, the boundary between the intermediate polars (with accretion disks) and polars (no accretion disks) is determined by the relation rm = rmin. if we calculate both the radii using the parameters of the ss cygni we see that this relation is satisfied at ba ≈ 106 g separating the two regimes of the flow. we should note that this estimate of magnetic field induction, separating the intermediate polars and polars, is more or less general for cvs, since the value of q there varies insignificantly. the typical flow structure in the polars is shown in fig. 3, where, like in figures 1 and 2, we plotted an iso-surface of density logarithm lgρ = −4.5 (in units of ρ(l1)) for the model 4 (10 7 g). in the models 4–7 no accretion disks form and material is accreted via funnel flows. in the models 4 and 5 the stream, starting at l1 point, splits into two flows when approaching the accretor’s surface. then the first flows falls onto the northern magnetic pole of the accretor, the second — onto the southern. with growing magnetic field the difference between the two poles becomes more obvious. in the model 6, for example, only one accretion flow occurs that then ends up at the southern pole. in the model 7 the magnetic field is so strong that it almost entirely controls the flow structure in the roche lobe of the accretor. the field captures material in the immediate vicinity of the inner lagrangian point and, splitting 62 on the influence of magnetic field on accretion processes in cvs it into two flows — more powerful southern and weaker northern, — transports it onto the accretor’s surface along the field lines. let us now consider how the described variations of the flow structure influence the accretion rate. in the upper-left panel of fig. 4 we plot the computed accretion rate ṁa as a function of ba. the vertical lines show specific values of the accretion rate variations. the main feature of this function is that it is non-monotonic. for the values ba < 10 6 g the accretion rate increases with increasing magnetic field and the amplitude of its variations decreases. at the point ba = 10 6 g the accretion rate approaches its maximal value. this can be explained by the fact that in the solutions with accretion disks the accretion rate is controlled by processes of angular momentum transfer. the growing magnetic field in intermediate polars makes magnetic breaking more effective and, hence, increases ṁa. after the point of maximum the accretion rate starts to decrease while ba continues to grow. as we have shown above starting at the point ba = 10 6 g no accretion disk forms in the system and the material is accreted via funnel flows. in this case the accretion rate is determined by the throughput of the stream. when the magnetic field grows the width of the stream decreases and the accretion rate falls. dependence of the accretion efficiency ṁa/ṁ on the induction of magnetic field is shown in the lower-left panel of fig. 4. with growing ba this value increases monotonically. in a transit range of magnetic field values (near ba ≈ 106 g) this dependence has an inflection point. the lowest effectiveness of accretion (30–40%) is approached for intermediate polars with weak magnetic fields. this may be explained by the fact that their disks are large and a portion of material may leave the accretor’s roche lobe and form a common envelope of the binary system [12]. in the polars (with strong magnetic fields) almost all the material, flowing from the l1 point onto the accretor, is accumulated and ”enclosed” in the stream. thus, in such systems the effectiveness of accretion reaches almost 100%. the non-monotonic shape of the ṁa(ba) curve, obtained in our calculations, does not contradict the monotonic character of the dependence of the accretion effectiveness on ba. the point is that the mass transfer rate ṁ also depends on the value of magnetic field. this dependence is shown in the upper-right diagram of fig. 4. in figure one can see that with growing ba the value of ṁ monotonically decreases. in a range of weak fields (ba < 10 6 g, intermediate polars) this decrease is almost invisible. however in strong fields (ba > 10 6 g, polars) the value of the mass transfer rate sharply drops (by almost an order of magnitude). this behavior may be explained by the fact that the strong magnetic field of the accretor generates extra-pressure in the vicinity of the l1 point, which prevents the outflow of the material from the donor’s envelope. thus, with growing ba the throughput of the accretion stream decreases. 3 influence of the asynchronous rotation of the white dwarf on the accretion rate the influence of the accretor’s spin-rotation on mhd flow structure in cvs can be characterized by a relation between the radius of the magnetosphere rm and the co-rotation radius rc that is a distance at which the velocity of rotation of the field lines is equal to the rotation velocity of material in the accretion disk. if the accretor rotates relatively slow (rc > rm) then the rotation velocity of the field lines at the boundary of the magnetosphere is lower than the keplerian. so the material, captured by the magnetic field can freely fall onto the accretor’s surface. this regime may be called the ”accretor” regime. when the accretor’s rotation is fast enough (rc < rm) a centrifugal barrier occurs at the boundary of the magnetosphere. this barrier prevents material from free falling onto the surface. this regime we may call the ”propeller”. in systems of ae aqr type the rotation of the white dwarf is so fast that no accretion disk can form. thus, for this regime we can use a term ”super-propeller” [5]. to investigate how the asynchronous rotation of the accretor influences the flow structure in a close binary system we conducted four runs of 3d numerical simulations for the period relation of pspin/porb: 0.1 (”accretor”); 0.033 (”equilibrium rotation”); 0.01 (”propeller”); and 0.001 (”super-propeller”) [5]. to preserve the generality we again use the parameters of the ss cygni. in all the runs the magnetic induction was set equal to 105 g and the magnetic axis is inclined by θ = 30◦. in the models ”accretor” and ”equilibrium rotation” the flow structure is analogous to the model 1, considered in previous section. in the ”propeller” regime, in the inner regions of the disk a magnetospheric cavity forms. its radius is 0.05–0.1a. at the boundary of this cavity the kelvin–helmholtz instability develops. since there is almost no material in the cavity, the accretion rate in the ”propeller” model is close to zero. on the other hand the mass transfer continues and the mass of the disk grows. in the systems with an appropriate combination of parameters (disk mass, value of magnetic field and rotation velocity) the material at the boundary of the cavity at a certain moment can push the magnetosphere down and initiate a short period of accretion. after the exceptional mass has been removed the system returns into the ”propeller” regime. in the ”super-propeller” regime the material, issuing from the inner lagrangian point, is captured by the 63 d. v. bisikalo, a. g. zhilkin rapidly rotating accretor’s magnetosphere, acquires additional angular momentum and gets ejected out of the accretor’s roche lobe. the accretion rate in this case is almost zero. in the lower-right panel of fig. 4 we show dependence of the accretion rate on the parameter porb/pspin that characterizes the spin velocity of the accretor. the data shown are for the case of ba = 10 5 g. in figure one can see that with the growing accretor’s spin velocity the accretion rate decreases. if the white dwarf rotates relatively slow, which corresponds to the ”accretor” regime, the accretion rate decrease is not significant in comparison with the case of synchronous rotation. however, when the spin velocity overcomes an equilibrium value the system enters the ”¡propeller” regime and the accretion rate sharply drops. in fact, in the ”propeller” and ”super-propeller” regimes the accretion rate is determined by the accretion of material from the common envelope but not of the material, issuing from the inner lagrangian point. 4 conclusions using the results of 3d numerical mhd simulations, we have considered the influence of the magnetic field value ba and the spin-rotation velocity of the white dwarf on accretion processes in cvs. analyzing the results of seven computational runs for models with ba equal to 105 g, 5 × 105 g, 106 g, 5 × 106 g, 107 g, 5 × 107 g and 108 g, we have found that the accretion rate ṁa is a non-monotonic function of ba. in the intermediate polars the growing ba makes the magnetic breaking more effective and, hence, increases ṁa. in the polars, where the accretion rate is determined by the throughput of the funnel flow the accretion rate decreases with growing ba. the maximal value of the accretion rate occurs in the systems at a stage of transit between the intermediate polars and polars. this value is ∼60% of the initially set mass transfer rate. we should note that unlike the absolute value of ṁa the accretion efficiency ṁa/ṁ monotonically increases with the increasing induction of magnetic field. in the polars it approaches the value of almost 100%. this is because the mass transfer rate decreases with growing ba, since strong magnetic field prevents the inflow of material into the accretor’s roche lobe through the inner lagrangian point. analyzing the results of four computational runs of 3d mhd simulations with various values of the period ratio (pspin/porb = 0.1, 0.033, 0.01 0.001) and fixed ba = 10 5 g we found that the accretion rate decreases with the increasing spin-rotation velocity. it is interesting to note that even in a system, living in the ”propeller” regime, the average accretion rate may be not equal to zero if the combination of such parameters as the disk mass, value of magnetic field and rotation velocity can cause quasi-periodic accretion bursts (falling of the portion of the disk mass onto the accreting star). acknowledgement this work was supported by the basic research program of the presidium of the russian academy of sciences, the russian foundation for basic research research (projects 11-02-00076, 12-02-00047, 13-02-00077, 13-02-00939), the federal targeted program ”science and science education for innovation in russia 2009– 2013”. references [1] warner, b. : 1995, cataclysmic variable stars, cambridge university press, cambridge. doi:10.1017/cbo9780511586491 [2] campbell, c.g. : 1997, magnetohydrodynamics in binary stars, kluwer acad., dordrecht. [3] norton, a.j., , wynn, j.a., somerscales, r.v. : 2004, apj 614, 349. doi:10.1086/423333 [4] zhilkin, a.g., bisikalo, d.v. : 2010, astron. rep. 54, 1063. [5] zhilkin, a.g., bisikalo, d.v., boyarchuk, a.a. : 2012, phys. usp. 55, 115. [6] giovannelli, f., gaudenzi, s., rossi, c., piccioni, a. p.j. : 1983, acta astronomica 33, 319. [7] zhilkin, a.g., bisikalo, d.v. : 2009, astron. rep. 53, 436. [8] zhilkin, a.g., bisikalo, d.v. : 2010, advances in space research 45, 437. doi:10.1016/j.asr.2009.09.006 [9] boyarchuk, a.a., et al. : 2002, mass transfer in close binary stars, taylor & francis, london. [10] fridman, a.m., bisikalo, d.v. : 2008, phys. usp. 51, 551. [11] lubow, s.h., shu, f.h. : 1975, apj 198, 383. doi:10.1086/153614 [12] sytov, a.yu., et al. : 2007, astron. rep. 51, 836. doi:10.1134/s1063772907100083 64 http://dx.doi.org/10.1017/cbo9780511586491 http://dx.doi.org/10.1086/423333 http://dx.doi.org/10.1016/j.asr.2009.09.006 http://dx.doi.org/10.1086/153614 http://dx.doi.org/10.1134/s1063772907100083 on the influence of magnetic field on accretion processes in cvs discussion christian knigge: you seemed to use ss cyg as a prototype of dn intermediate polar, out what is the empirical evidence for this? dmitry bisikalo: we just used ss cyg as a dn illustrative example. the parameters of the ss cyg (components masses, radii, separation, etc.) are quite typical for intermediate polars, so we can consider these numerical results as a good illustration of the flow structure in the intermediate polars. giora shaviv: if i saw right, you do not have radiative transfer, only local energy sinks. how it works when the disk is optically thick? dmitry bisikalo:yes, this approach is strictly applicable for optically thin disks. if you consider optically thick disks it is necessary to add the radiative transfer in some form to the model, and we are working on this. fortunately, we can use the assumption of optically thin disks for a significant sample of cvs. 65 introduction what happens with the flow structure if ba grows influence of the asynchronous rotation of the white dwarf on the accretion rate conclusions acta polytechnica ctu proceedings doi:10.14311/app.2015.1.0045 acta polytechnica ctu proceedings 2:45–50, 2015 © czech technical university in prague, 2015 available online at http://ojs.cvut.cz/ojs/index.php/app impact assessment of image feature extractors on the performance of slam systems taihú pirea, ∗, thomas fischera, jan faiglb a university of buenos aires, intendente güiraldes 2160, ciudad autónoma de buenos aires, argentina b department of computer science, faculty of electrical engineering, czech technical university in prague, technická 2, 166 27, prague, czech republic ∗ corresponding author: tpire@dc.uba.ar abstract. this work evaluates an impact of image feature extractors on the performance of a visual slam method in terms of pose accuracy and computational requirements. in particular, the s-ptam (stereo parallel tracking and mapping) method is considered as the visual slam framework for which both the feature detector and feature descriptor are parametrized. the evaluation was performed with a standard dataset with ground-truth information and six feature detectors and four descriptors. the presented results indicate that the combination of the gftt detector and the brief descriptor provides the best trade-off between the localization precision and computational requirements among the evaluated combinations of the detectors and descriptors. keywords: image features, visual slam, stereo vision. 1. introduction during the last decade, the simultaneous localization and mapping (slam) problem has been one of the main research interests in mobile robotics. particularly, the use of cameras as the main sensors has been given a special attention [1] [2] [3] [4] because of their benefits such as low-cost and passive sensing. in vision-based slam approaches, local image features are used to build a map and simultaneously estimate the robot pose using the environment landmarks represented as the image features. in this way, the map is represented as a sparse point cloud, where each point results from triangulating salient points (image features) matched from a pair of stereo images. currently, there exist several local image feature extractors in the literature. a feature extractor is a combination of a salient point (called keypoint) detection procedure and a computation of a unique signature (called descriptor) for each such a detected point. the most commonly used detectors are sift [5], surf [6], star [7], gftt [8], fast [9], and relatively recently proposed orb [10], while among the most used descriptors we can mention sift, surf, orb, brief [11], and brisk [12]. in visual slam systems, the feature extraction process has a huge impact on the accuracy of the whole system. on one hand, the precision of the robot localization is heavily correlated to the sparsity of features in images and the ability to track them for a long period during the robot navigation, even from different points of view. on the other hand, if the number of points in the map grows too quickly, it may slow down the whole system. to be able to keep the response of the system under real-time constraints, images have to be dropped or other parts of the system, like optimization routines, need lower computational requirements. in this work, we evaluate the impact of different state-of-the-art feature extractors on the performance of the visual slam localization method. in particular, the evaluation is based on the stereo visual slam approach s-ptam introduced in [4]. the presented results indicated that the combination of the gftt detector and brief descriptor is the most reliable choice for our slam system among the other evaluated combinations. the rest of the paper is organized as follows. section 2 presents overview of the related work while section 3 summarizes the most used feature detectors and descriptors in the visual slam literature. section 4 briefly comments the considered stereo visual slam system using for the evaluation. in section 5, we present the evaluation of the features extractors and the achieved results. section 6 is dedicated to the conclusions and future work. 2. related work several evaluations of features extractors can be found in literature. each of them is driven by the particular application or issue at the hand they are aimed to address. for example, in [13], authors evaluate several features extractors in the context of autonomous navigation in outdoor environments under seasonal changes. they came to a conclusion that the best performing method is the star–brief combination of the detector–descriptor, which outperforms sift by more than thirty percentage points. in addition, they argued that the star–brief extractor is also less computationally demanding than other extractors and thus it seems to be the most suitable feature 45 http://dx.doi.org/10.14311/app.2015.1.0045 http://ojs.cvut.cz/ojs/index.php/app t. pire, t. fischer, j. faigl acta polytechnica ctu proceedings detector–descriptor for navigational purposes. on the other hand, authors of [14] provide a performance comparison of feature extractors against illumination changes in outdoor scenes in the context of the visual navigation. they concluded that the configuration of the fast–surf is the optimal in their setup. besides, they report that this combination provides an effective computational time per image, which is favorable for the real-time vision-based navigation application. the work [15] compares contemporary point features detector and descriptor pairs in order to determine the best combination for the robot visual navigation. the authors concluded that the fast– brief combination is a good choice when processing speed is an important parameter of the system setup. they also argued that under camera movement conditions, additional computational cost—needed for the descriptors and detectors that are robust to in-plane rotation and large scaling— seems to be unjustified. however, they do not tested the method in a real slam application. regarding the aforementioned evaluation of the detectors and descriptors, the work presented in this paper is within the context of the full 6dof slam. 3. local image features an image feature extractor consists of detection and description phases. the feature detector serves to locate salient areas of the image while the feature descriptor captures information about the local neighborhood of the detected area. here, we provide a brief overview of the considered feature extractor and descriptor algorithms in this evaluation study. sift – scale invariant feature transform [5]. an established feature detector with a high precision and good robustness, which is known to be computationally demanding. surf – speeded up robust features [6] is a similar to sift, but it is computationally less demanding due to approximations. star – a modified version of the censure (center surrounded extrema) [7] detector, which is computationally less demanding at the expense of a lower precision. brief – binary robust independent elementary features [11] is a descriptor that describes an image area using a number of intensity comparisons of random pixel pairs. it is saved as a binary string, which reduces the computational complexity of the subsequent matching. fast – features from accelerated segment test [9] is a feature detector focused on lowering the computational cost. brisk – binary robust invariant scalable keypoints [12] is a scale and rotation invariant version of brief, but unlike brief, it uses a deterministic comparison pattern. orb – oriented fast and rotated brief [10] is another attempt to achieve a scale and rotation invariant brief, as a computationally efficient alternative to sift and surf. it uses the fast detector to achieve low computational requirements. gftt – a detector focused on selecting features relevant to motion tracking by analyzing the amount of information they provide for that particular task [8]. the surf and sift descriptors rely on the their own detectors, which are also considered in the presented evaluation. however, for the brief and brisk binary descriptors the considered detectors are the gftt, fast and star which results in the additional six combinations of the detector–descriptor pairs in the presented evaluation. 4. overview of s-ptam s-ptam [4] is a stereo visual slam method for a large scale map navigation based on the monocular parallel tracking and mapping (ptam) method introduced in [1]. the method consists of two processes working in parallel: 1) the tracking of the detected features and; 2) creating a map of the features (mapping). during a robot navigation, the method works as follows. s-ptam extracts features from the incoming stereo images to match and construct a virtual map of the environment. the newly extracted feature descriptors are matched against descriptors of the points stored in the map according to the estimated field of view. the matches may then be used to refine the estimated camera pose using an iterative least squares minimization method, e.g., using the levenberg-marquardt algorithm. the particular stereo matches between the features that cannot be matched to the map are triangulated and inserted as new map points, for the tracking of future frames. in parallel, a map refinement algorithm is running. it is also based on the levenberg-marquardt optimization that continuously performs the bundle adjustment on the current local portion of the map. in [4], s-ptam uses the gftt feature detector and the brief descriptor extractor. in this work, we consider other combinations of the detector–descriptor to evaluate an impact of the combination to the performance of the localization and mapping processes. 5. evaluation the kitti vision benchmark suite [16] is used to evaluate s-ptam for each type of considered detector– descriptor configuration. in particular, we present the results obtained for the sequence 00, shown in figure 1. the sequence records the stereo camera frames captured by a moving car in an urban scenario for almost 4 km long path. the particular parameters of 46 vol. 2/2015 impact assessment of image feature extractors on slam the evaluated feature extractors have been selected in such a way that allows s-ptam to run without ever loosing localization. they have been tuned from a strong restrictive value and then relaxed until the method completes the whole sequence. the parameters are listed in table 1. detector / parameter valuedescriptor sift noctavelayers 1 l2normthreshold 100 surf hessianthreshold 1000 noctaves 1 l2normthreshold 0.2 star responsethreshold 20 brief bytes 32 hammingthreshold 25 fast threshold 60 brisk hammingthreshold 100 orb nfeatures 2000 nlevels 1 hammingthreshold 50 gftt nfeatures 2000 mindistance 15.0 table 1. parameters used for feature detectors and descriptors. the parameters which do not appear in the list use the default value in the opencv implementation. in the case of the binary descriptors, the hamming distance is used to compute the valid matches while the l2 norm is used for the surf and sift descriptors. the evaluation has been performed using an intel core i7 processor with 4 cores running at 2.2 ghz. although s-ptam strongly exploits parallelism, the experiments were run in a sequential fashion that allow us to simulate ideal conditions and abstract from the limitations of the available computational power. this ensures that no frames are dropped and that the iterative optimization routines always converge or reach a maximum threshold of iterations. nevertheless, the tracking process performs pose optimization using an iterative algorithm; so, the less time is used in the features extraction, the more iterations the method can compute. figure 2 shows a characterization of the total tracking time for each pair of frames, as achieved by using the evaluated extractors. moreover, the iterative least-squares optimization, which is utilized in the mapping and tracking processes, depends linearly on the number of tracked points (the density of the map). thus, regarding the computational burden, the map should be as small as possible while the map points should contain strong enough features to support a robust tracking of the figure 1. path tracked by every method run under different extractors, against the ground truth. the path is nearly 4 km long. the shown distances at the axes are in meters. figure 2. total tracking time achieved by each configuration frames. table 2 shows the final number of points contained in the map after finishing each trial for a particular combination of feature detector and descriptor. in figure 3, we can see how the map size impacts directly on the temporal performance of the tracking process. combinations of the detector–descriptor that build the most dense maps also take the longest time to compute. differences in the map size for the evaluated descriptor in the feature extractors with the same detector can have two reasons. the first reason is that new 47 t. pire, t. fischer, j. faigl acta polytechnica ctu proceedings figure 3. tracking time without taking into account feature extraction points are created from the stereo features only if these features are not matched to the map. the second reason is that the points marked as outliers during the refinement processes are discarded. in the first case, this can be caused by descriptors that are not robust enough to be matched to the map for a long time. in the second case, the descriptor matching may be too permissive and it allows bad matches that are later discarded as outliers. extractor final map size gftt / brief 990 455 gftt / brisk 1 314 356 sift / sift 1 581 876 star / brief 1 893 372 surf / surf 2 059 879 fast / brief 2 420 652 star / brisk 2 447 418 fast / brisk 3 207 003 orb / orb 5 192 885 table 2. the number of points contained in the map after completing the sequence for each evaluated extractor, in ascending order. since the goal of this work is to assess the impact of the feature extractor choice also on the accuracy of the slam method, the achieved performance is presented as two independent relative errors for each estimated pose: �t for the translation error and; �θ for the orientation. let xk be the estimated pose at the frame k, which can be decomposed as the translation tk and the rotation rk. let x∗k be the reference pose, which can be decomposed in the same fashion. the aforementioned errors are computed as �t,k+1 = ‖(tk tk+1) ( t∗k t ∗ k+1 ) ‖, �θ,k+1 = angle ( (rk rk+1) ( r∗k r ∗ k+1 )) , where is the inverse of the standard motion composition operator. for pure translations, we can rewrite t1 t2 = t2 −t1, and for the pure rotations as r1 r2 = rt1r2. ‖x‖ stands for the euclidean norm and angle (r) extracts the magnitude of the rotation. the computed errors are shown in figure 4 and figure 5, respectively. although the angular deviation to the ground truth, shown in figure 5, seems to be similar in all methods, the same is not true for the translation error, as it can be seen in figure 4. the brisk descriptor seems to be a more reliable with the fast detector, while the same holds for the brief descriptor with the gftt detector. figure 4. relative translation error figure 5. relative orientation error for completion, the absolute errors �′t,k = ‖tk t∗k‖ �′θ,k = angle (rk r∗k) are shown in figure 6 and figure 7. 48 vol. 2/2015 impact assessment of image feature extractors on slam figure 6. absolute translation error figure 7. absolute orientation error 6. conclusions in this paper, we present an evaluation of the impact of different state-of-the-art image feature extractors on the performance of the slam method proposed in [4]. the kitti benchmark suite dataset with a ground truth is used to evaluate the achievable precision of the method for different feature extractors. based on the presented results, the main conclusion is that the gftt detector is the most suitable choice for the best performance in the evaluated dataset. the gftt (accompanied with the brief or brisk descriptors) outperforms the other methods in terms of the required computational time and the map quality. although the map density is far smaller, the computed translation error is similar, even slightly better, than the one achieved by other extractors. this insight can be interpreted as the most useful features (regarding the navigation) are extracted while the descriptor also support efficient matching resulting in a more precise localization. recently, a novel stereo feature extractors have been proposed, e.g., [17], which motivates us to consider them in s-ptam. an evaluation of the novel extractors is a subject of our future work. acknowledgements this work is a direct result of the bilateral cooperation program between the czech and argentinian republics support by the argentinian project arc/14/06 and travel support of the czech ministry of education under the project no. 7amb15ar029. the work of j. faigl is supported by the czech science foundation (gačr) under the research project no. gj15-09600y. references [1] g. klein, d. murray. parallel tracking and mapping for small ar workspaces. in ismar, pp. 1–10. ieee computer society, washington, dc, usa, 2007. doi:10.1109/ismar.2007.4538852. 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[17] r. arroyo, p. alcantarilla, l. bergasa, et al. fast and effective visual place recognition using binary codes and disparity information. in iros, pp. 3089–3094. 2014. doi:10.1109/iros.2014.6942989. 50 http://dx.doi.org/10.1109/acpr.2013.159 http://dx.doi.org/10.1177/0278364913491297 http://dx.doi.org/10.1109/iros.2014.6942989 acta polytechnica ctu proceedings 2:45–50, 2015 1 introduction 2 related work 3 local image features 4 overview of s-ptam 5 evaluation 6 conclusions acknowledgements references 35 acta polytechnica ctu proceedings 2(1): 35–40, 2015 35 doi: 10.14311/app.2015.02.0035 white dwarfs in cataclysmic variables: an update e. m. sion1, p. godon1 1department of astrophysics & space science, villanova university, 800 lancaster ave., villanova, pa 19085, usa corresponding author: edward.sion@villanova.edu abstract in this review, we summarize what is currently known about the surface temperatures of accreting white dwarfs in nonmagnetic and magnetic cataclysmic variables (cvs) based upon synthetic spectral analyses of far ultraviolet data. we focus only on white dwarf surface temperatures, since in the area of chemical abundances, rotation rates, wd masses and accretion rates, relatively little has changed since our last review, pending the results of a large hst go program involving 48 cvs of different cv types. the surface temperature of the white dwarf in ss cygni is re-examined in the light of its revised distance. we also discuss new hst spectra of the recurrent nova t pyxidis as it transitioned into quiescence following its april 2011 nova outburst. keywords: cataclysmic variables dwarf novae intermediate polars spectroscopy uv individuals: t pyx, cc syg. 1 introductory overview the white dwarfs in cataclysmic variables (hereafter cvwd) are the central engines of the observed outbursts, either as potential wells for the release of gravitational energy during accretion (dwarf nova dn), or as the sites of explosive thermonuclear runaway (tnr) shell burning (classical novae), steady shell burning (supersoft x-ray binaries) or instantaneous collapse and total thermonuclear detonation if the wd reaches the chandrasekhar limit (type ia supernova? sn ia). therefore, the accreting wds serve as probes of explosive evolution and accretion physics and diffusion, as they bear the thermal, chemical and rotational imprint of their long term accretion and thermonuclear history (sion 1991 & 1995; townsley & bildsten 2003; townsley & gänsicke 2009). deeper physical insights however require a larger number of chemical abundances, rotation rates, surface temperatures, mass accretion rates, and masses for each spectroscopic subclass of cvs. only then can any definitive conclusions be drawn. in order to adequately sample the parameter space (mwd, i, ṁ, teff , porb) of the dns, nova-like variables, and magnetic cvs, a large go program was approved in cycle 20 (b.gänsicke, principal investigator) to secure high quality cos spectra for cv classes underrepresented in the current overall cv sample. the data is in hand and undergoing analysis at the time of this writing. ultimately, we hope to obtain data for > 30 cvs per class. therefore, in this review our focus is restricted to white dwarf surface temperatures, since in the area of chemical abundances, rotation rates, wd masses and accretion rates, relatively little has changed since our last review in the 2011 palermo meeting proceedings. in section 2, our fuv analysis techniques are briefly summarized, section 3 we address how the revised (shorter) distance to ss cygni affects the results of our analysis of the fuse + hst stis spectra of ss cygni by sion et al. (2010), in section 4 we tabulate and display the current distribution of cvwd surface temperatures versus orbital period and in the final section, we include some remarks on new hst spectra of the recurrent nova t pyxidis. 2 synthetic spectral analysis of fuv spectra of cvwds we have modeled the fuv spectra of wds in cvs during dn quiescence and nova-like low states from iue, fuse and hst (fos, ghrs, stis, cos) archival data, and through our ongoing collaboration with past hst surveys led by b.gänsicke, and p. szkody. the iue archival spectra and high quality fuse and hst fuv data are fit with the latest versions of the tlusty/synspec model photosphere code and model accretion disk codes (hubeny 1988; hubeny & lanz 1995). we are taking into account the bl explicitly in our modeling of the fuv spectra of disk accreting systems, by replacing the very inner rings of the standard accretion disk model with high temperature rings in agreement with the temperature and density in the boundary layer. this improves the model fitting at the shorter wavelengths. as an example of our cvwd 35 http://dx.doi.org/10.14311/app.2015.02.0035 e. m. sion, p. godon photosphere fitting, in fig.1, we display a wd solar composition fit to the hst stis spectrum of v442 cen with e(b-v) = 0.10. the wd model has t=47,000k ±2000k and log(g) = 8.3±0.2, vsin(i) = 300km/s ±50 km/s (sion et al. 2008). figure 1: the best-fitting single temperature wd fit to the hst/stis spectrum of v442 cen. the model consists of a 47,000k wd with log(g)=8.3 and scaled to a distance of 328pc. 3 the white dwarf in ss cygni:the vlbi and corrected hubble fgs distance schreiber & lasota (2007), and references therein, pointed out that the previously published hubble fgs parallax of ss cygni, with a distance of 166 +/12 pc, posed serious problems for the disc instability model (dim) of dwarf nova outbursts because it would fail to explain the absolute magnitude during outburst. with the new vlbi-derived distance (miller-jones et al. 2013) to ss cygni (114 pc), and the corrected hst distance (nelan & bond 2013) of 120 pc instead of 166 pc, the concern for the validity of dim is alleviated. an obvious question is: how does the new, shorter distance affect the analysis of sion et al. (2010) regarding their detection of the wd during quiescence and their derivation of the wd’s surface temperature from the modeling of fuse + hst stis spectra? to answer this question, we have carried out model fits using the new distance in the range of 114 pc to 120 pc. when we used single, steady state disk models, we obtained the same results as in sion et al. (2010). the disk does not fit the flux in the shorter wavelengths unless the accretion rate is very large, which is inconsistent with dwarf nova quiescence, and the model-derived distance is far too large. when we combined a model disk with a model white dwarf photosphere (using the bitner et al. wd mass as in sion et al. (2010), the correct distance to ss cygni was obtained for a low accretion rate and a lower temperature white dwarf. however, the fits with the wd + disk are inferior to the fits where the wd dominates the fuv flux. in order to match the best fit solution to the fuse + hst stis wavelength range in sion et al. (2010), the white dwarf must have teff = 45000k 50000k for a distance of 112 pc to 120 pc and a white dwarf mass mwd = 0.95m�. this value of the wd mass is within the error range of the bitner et al. mass. our conclusion about the teff of the wd in ss cygni during quiescence is essentially unchanged from sion et al. (2010). 4 the current distribution of cv white dwarf surface temperatures versus orbital period the effective temperatures teff of the wds (obtained during dwarf nova quiescence and nova-like low states when the wd is exposed), are the critical key to revealing the thermal response of the wd to mass accretion and the long term accretion rate < ṁ > which, through compressional heating by the accreted material, is linked to the wd surface temperature (sion 1995; townsley & bildsten 2003; townsley & gänsicke 2009). in table 1, we have tabulated what we regard as the most reliably secured wd surface temperatures. the first column gives the name of the cv, the second column the cv subtype, the third column, the orbital period in minutes, the fifth column the wd teff and the sixth column, the reference for the temperature. these temperatures are derived from a variety fuv spectra and all are known with a precision of at least 3000k and in the majority of cases, better that +/2000k. with the paucity of fuv-derived temperatures per cv subclass, one should not ignore usable fuv data from iue, fuse and hut because the ”quality” of the data is deemed inferior to hst stis and cos. the current distribution of wd teff against the orbital period porb is displayed in fig.2. (see also fig.4 in townsley & gänsicke 2009). note the continued relatively sparse coverage in temperature of the wds in cvs above the period gap (porb > 3hr), compared with the coverage below the period gap. it is possible that the apparent trend toward higher temperatures with increasing porb (i.e. higher long term average < ṁ >) could be due to observational selection since the wd teff ’s were derived primarily in the fuv where the planckian peak occurs for hotter accreting wds. for example, copperwheat et al. (2010) derive a teff = 10 − 15, 000k in the optical for the wd in the eclipsing dn ip peg which has porb = 3.8 h while the eclipsing system sdss1006 with porb = 4.46 h, has teff = 16, 500 ± 2000k (southworth et al. 2009) in the optical. both of these objects should contain much hotter wds (higher < ṁ >) for their orbital periods. 36 white dwarfs in cataclysmic variables: an update table 1: the temperature of cv white dwarfs system cv type porb (h) teff (k) ref sdss1507 dn/su 1.11 11000 szkody et al. (2010a) gw lib dn/wz 1.280 14700 szkody et al. (2002a) bw scl dn? 1.304 14800 gänsicke et al. (2005) ll and dn/wz 1.321 14300 howell et al. (2002) pq and dn/su 1.34 12000 szkody et al. (2010a) ef eri am 1.350 9500 szkody et al. (2010b) sdss j1610-0102 dn? 1.34 14500 szkody et al. (2007) v455 and dn/su 1.35 10500 szkody et al. (2013) hs2331+3905 dn 1.351 10500 araujob-betancor et al. (2005a) al com dn/wz 1.361 16300 szkody et al. (2003) wz sge dn/wz 1.361 14900 sion et al. (1995) sw uma dn/su 1.364 13900 gänsicke et al. (2005) sdss0919 dn 1.36 13500 szkody (2014) sdss1035 dn? 1.37 10500 southworth et al. (2006); littlefair et al. (2006b) hv vir dn/wz 1.370 13300 szkody et al. (2002b) sdss1339 dn 1.38 12500 szkody et al. (2010a) sdss2205 dn 1.38 15000 szkody et al. (2010a) wx cet dn/wz 1.399 13500 sion et al. (2003) t leo dn/su 1.41 16000 hamilton & sion (2004) eg cnc dn/wz 1.44 12300 southworth et al. (2006) xz eri dn/su 1.468 15000 szkody et al. (2010a) sdss1514 dn 1.48 10000 szkody (2014) dp leo am 1.497 13500 schwope et al. (2002) v347 pav am 1.501 11800 araujob-betancor et al. (2005b) bc uma dn/su 1.503 15200 gänsicke et al. (2005) ek tra dn/su 1.509 18000 godon et al. (2008) vy aqr dn/wz 1.514 14500 sion et al. (2003) oy car dn/su 1.515 15000 cheng et al. (2000) sdss0131 dn/su 1.63 14500 szkody et al. (2010a) vv pup am 1.674 11900 araujob-betancor et al. (2005b) v834 cen am 1.692 14300 araujob-betancor et al. (2005b) ht cas dn/su 1.768 14000 wood et al. (1992) vw hyi dn/su 1.783 22000 godon et al. (2008) cu vel dn/su 1.88 18500 gänsicke & koester (1999) mr ser am 1.891 14200 araujob-betancor et al. (2005b) bl hyi am 1.894 13300 araujob-betancor et al. (2005b) st lmi am 1.898 10800 araujob-betancor et al. (2005b) ar uma am 1.93 20000 schmidt et al. (2005) rej1225 dn/su 1.99 12000 szkody et al. (2010a) ef peg dn/wz 2.01 16600 howell et al. (2002) dv uma dn/su 2.138 20000 feline et al. (2004) hu aqr am 2.084 14000 gänsicke (1999) qs tel am 2.332 17500 rosen et al. (2001) sdss j1702+3229 dn/su 2.402 17000 littlefair et al. (2006a) tu men dn 2.813 28000 sion et al. (2008) am her am 3.094 19800 gänsicke et al. (1995) mv lyr nl/vy 3.176 45000 godon et al. (2012) dw uma nl/vy 3.279 50000 araujob-betancor et al. (2003) tt ari nl/vy 3.301 39000 gänsicke et al. (1999) ip peg dn 3.80 15000 southworth et al. (2009) vy scl nl 3.99 45000 hamilton & sion (2008) v1043 cen am 4.190 15000 araujob-betancor et al. (2005a) ww cet dn 4.220 26000 godon et al. (2008) ugem dn/ug 4.246 30000 sion et al. (2001) ssaur dn/ug 4.391 34000 godon et al. (2012) sdss1006 dn 4.46 16500 southworth et al. (2009) v895 cen am 4.765 14000 araujob-betancor et al. (2005b) rx and dn/zc 5.037 34000 sion et al. (2001) ss cyg dn/ug 6.60 47000 sion et al. (2010) vy scl nl 3.99 45000 hamilton & sion (2008) em cyg dn/zc 6.98 40000 godon et al. (2012) tt crt dn 7.30 29000 sion et al. (2008) ru peg dn 8.99 70000 godon et al. (2012) v442 cen dn 11.04 47000 sion et al. (2008) 37 e. m. sion, p. godon figure 2: effective white dwarf temperature as a function of the orbital period. the references for the indivdual temperatures can be found in sion et al. 2008, townsley & gänsicke 2009, araujo-betancor 2005a & b and references therein). the traditional magnetic braking above the period gap is shown between the parallel diagonal solid lines. on the right hand side are the time-averaged accretion rates correpsonding to the temperature scale on the left hand side of the diagram. shown for comparison between the dotted lines is the long term evolutionary path of a 0.8 solar mass white dwarf (with an initial core temperature of 30 million degrees k) which has undergone 1000 nova outburst cycles accreting at the long term rate of 10−8m�/yr 5 the hst cos + stis spectra of the recurrent nova t pyxidis:a progress report the interstellar reddening e(b-v) of the recurrent nova t pyxidis is a critical parameter in the determination of the best-fitting model parameters. the uv spectra exhibit a minimum near 2175 a which is due to the interstellar extinction. since the galex spectrum is the most reliable (i.e. highest s/n ratio) in that wavelength region, we used the galex spectrum to determine e(b-v). the value of e(b-v) for which the 2175 a feature disappears from the dereddened spectrum, e(b-v)=0.35, is taken as the e(bv) value towards t pyx (see fig.3). we use this value to deredden the iue, galex and hst spectra, and we also consider the effects of different reddening values on our results. the hst stis and cos spectra obtained in december 2012 and july 2013 are identical. thus, we co-added them to improve the signal-to-noise (s/n). we have found that the pre-outburst iue and galex spectra together with the hst post-outburst spectrum. we note that the uv flux has remained constant not only before the outburst, but it has now come back precisely to the same value. this is an indication that 1600 1800 2000 2200 2400 2600 2800 1e-14 1e-13 1e-12 e(b-v) 0.00 0.10 0.20 0.25 0.30 0.35 0.40 0.45 0.50 a log (flux) er g/ s/ cm ^2 /ao o figure 3: the merged galex-iue spectrum of t pyx has been dereddened for different values of e(b-v) as indicated on the right. the 2175å feature associated with the reddening is clearly seen in absoprtion for low balues of e(b-v), and it appears as extra flux for large values of e(b-v). we deduce that the reddening towards t pyx must be e(b-v)=0.35, the value for which the 2175å feature vanishes. 38 white dwarfs in cataclysmic variables: an update the mass accretion rate remained constant before and after the outburst (see godon et al. 2014). 0.4 0.6 0.8 1 1.2 1.4 1.6 1e-06 1e-05 0.0001 0.001 mwd/msun envelope mass (msun) e(b-v)=0.45 luv e(b-v)=0.35 luv e(b-v)=0.25 luv e(b-v)=0.50 disk e(b-v)=0.35 disk [1] e(b-v)=0.50 e(b-v)=0.25 disk [2] accreted accreted accreted accreted accreted accreted ejected ejected extrapolation figure 4: the mass of the accreted envelope and ejected envelope are shown as a function of the wd mass for different values of the reddening for t pyx. our disk model results are drawn with a solid line (“disk”). the lower limit of the accreted envelope inferred from the uv flux is shown (dotted line, “luv”). in comparison we show the ejected envelope [1] as estimated by nelson et al. (2012) (square symbol), as well as [2] computed in patterson et al. (2013) (dashed line). the accreted envelope is larger than the ejected envelope, except for a value of e(b-v)=0.25 combined with a wd mass ∼ 0.9m� and larger. however, the best results were obtained for a larger value of the reddening e(b − v ) > 0.30. using the value of the reddening we derived, together with the new distance estimate of 4.8 kpc (sokoloski et al. 2013), we fit the observed (back-toquiescence) hst spectra with disk models for different wd masses. we then computed the accreted mass over a period of 44yrs, which we then compared to the estimates of the ejected envelope mass (during the 2011 outburst). we recapitulate our results in fig.4, where we also consider different reddening values for the sake of completness. our main finding is that the mass of the accreted material between the last two outbursts is larger than the mass of the ejected envelope in the last outbursts, unless the reddening is e(b-v)=0.25 (or smaller) and the mass of the wd is ∼ 0.9m� or larger. 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communication [50] townsley, d.,& bildsten, l.: 2003, apj, 596, l227 [51] townsley, d., & gänsicke, b.: 2009, apj, 693, 1007 doi:10.1088/0004-637x/693/1/1007 [52] wood, j. h., horne, k., & vennes, s.: 1992, apj, 385, 294 40 http://dx.doi.org/10.1046/j.1365-8711.2001.04141.x http://dx.doi.org/10.1086/431969 http://dx.doi.org/10.1086/586699 http://dx.doi.org/10.1088/2041-8205/716/2/l157 http://dx.doi.org/10.1088/2041-8205/770/2/l33 http://dx.doi.org/10.1111/j.1365-2966.2006.11042.x http://dx.doi.org/10.1086/511854 http://dx.doi.org/10.1088/0004-637x/716/2/1531 http://dx.doi.org/10.1088/0004-637x/775/1/66 http://dx.doi.org/10.1088/0004-637x/693/1/1007 introductory overview synthetic spectral analysis of fuv spectra of cvwds the white dwarf in ss cygni:the vlbi and corrected hubble fgs distance the current distribution of cv white dwarf surface temperatures versus orbital period the hst cos + stis spectra of the recurrent nova t pyxidis:a progress report 151 acta polytechnica ctu proceedings 1(1): 151–156, 2014 151 doi: 10.14311/app.2014.01.0151 argo-ybj: highlights and prospects giuseppe di sciascio1 on behalf of the argo-ybj collaboration 1infn sezione di roma tor vergata, viale della ricerca scientifica 1, i-00133 roma corresponding author: disciascio@roma2.infn.it abstract the argo-ybj experiment has been in stable data taking for 5 years at the yangbajing cosmic ray laboratory (tibet, p.r. china, 4300 m a.s.l., 606 g/cm2). with a duty-cycle greater than 86% the detector collected about 5·1011 events in a wide energy range, from few hundreds gev up to the pev. a number of open problems in cosmic ray physics has been faced exploiting different analyses. in this paper we summarize the last results in gamma-ray astronomy and in the cosmic ray physics and introduce the lhaaso project, mainly driven by the chinese community, to study the cosmic ray physics up to 1017 ev. keywords: cosmic rays extensive air showers gamma-ray astronomy argo-ybj. 1 the argo-ybj experiment argo-ybj is a full coverage air shower detector located at the yangbajing cosmic ray laboratory (4300 m a.s.l., 606 g/cm2, tibet, pr china) devoted to the study of gamma rays and cosmic rays. exploiting the high altitude and the full coverage technique, argoybj can detect gamma rays with an energy threshold as low as a few hundred gev. the detector consists of a ∼74×78 m2 carpet made of a single layer of resistive plate chambers (rpcs) with ∼93% of active area, surrounded by a partially instrumented (∼20%) area up to ∼100×110 m2. the detector has a modular structure with a high granularity, that provides a detailed view of the shower front. the smallest space-time unit, called “pad”, has a size of 55.6×61.8 cm2. the time and the location of each fired pad are recorded and used to reconstruct the position of the shower core and the arrival direction of the primary particle. the point spread function (psf), the pointing accuracy and the energy calibration of the detector have been evaluated using the moon shadow technique, i.e. the deficit of cosmic rays in the moon direction. for a detailed description of the detector performance see (aielli et al., 2006, aielli et al. 2009, bartoli et al., 2011). since november 2007 to january 2013 the argoybj experiment monitored with high duty cycle (∼86%) the northern sky at tev photon energies. with a cumulative sensitivity ranging from 0.24 to ∼1 crab units, depending on the declination, six sources have been observed with a statistical significance greater than 5 standard deviations (s.d.) in the declination band from -10◦ to 70◦. in this paper the last results obtained in gamma-ray astronomy and in the study of the anisotropy in the cr arrival direction distribution are summarized. 2 northern sky survey the argo-ybj data used in this analysis were collected from november 2007 to january 2013, with a total observation time of 1670.45 days. the total number of events selected with a zenith angle less than 50◦ is about 3×1011. they are used to fill a map in celestial coordinates (right ascension and declination) with 0.1◦×0.1◦ bins, covering the declination band from -10◦ to 70◦. )σ significance ( -5 0 5 10 15 20 n u m b e r o f e n tr ie s p e r b in 1 10 210 3 10 410 5 10 6 10 all region °<l<90°,20°|b|<2 figure 1: significance distribution of the whole sky map (thick solid line). the thin solid line represents the best gaussian fit. the significance distribution of the galactic plane region with | b |< 2◦ and 20◦ < l < 90◦ is shown by the thick dotted line. the thin dotted line represents the best gaussian fit for the same region. 151 http://dx.doi.org/10.14311/app.2014.01.0151 giuseppe di sciascio figure 2: significance map of the northern sky as seen by the argo-ybj experiment in vhe γ-ray band. the two dotted lines indicate the galactic latitudes b = ±5◦. galactic longitude (deg) g al ac tic la tit ud e (d eg ) -10 -5 0 5 10 -4 -2 0 2 4 6 °dec=50 °dec=40 °dec=30 °dec=20 °dec=10 °dec=0 2030405060708090 figure 3: significance map of the galactic plane region with | b |< 10◦ and 20◦ < l < 90◦ obtained by the argo-ybj experiment. the circles indicate the position of all the known vhe sources. the open stars mark the location of the gev sources of the second fermi-lat catalog. the significance distribution of the whole map bins is shown in fig.1. the distribution, with a mean value of 0.002 and σ=1.02, closely follows a standard gaussian distribution except for a tail with large positive values, due to the excesses from several gamma ray emission regions, shown in fig. 2. table 1 lists the locations of all the regions with pre-trial significance greater than 4.5 s.d. . for each independent region, only the coordinates of the pixel with the highest significance are given. based on the distribution of negative points (fig. 1), a significance threshold of 4.5 s.d. corresponds to ∼ 2 false sources in our catalog. the significance distribution of the inner galactic plane region (longitude 20◦ < l < 90◦ and latitude | b |< 2◦) is also shown in fig. 1. the gaussian fit of the distribution has a mean value of 0.40 and σ=1.04. also in this case a tail of significant excesses is present. the locations of all the excesses with a significance greater than 4.0 s.d. are listed in table 1. fig. 3 shows the inner galactic plane region in galactic coordinates. since the smoothing radius is larger than the bin width, the significances in adjacent bins are correlated, and a monte carlo simulation is necessary to correctly evaluate the post-trial probabilities. according to our simulations, the probability to have at least one source with a pre-trial significance > 5.1 σ (> 4.0 σ) anywhere in the map (in the inner galactic plane) due to background fluctuations is 5%. however, since in the sky region monitored by argo-ybj only ∼70 known vhe emitters exist, the post-trial significance increases for the candidate sources associated to a counterpart. detail about different sources are discussed in chen s. et al. (2013). 152 argo-ybj: highlights and prospects table 1: location of the excess regions argo-ybj ra dec s associated name (deg) (deg) (s.d.) tev source j0409−0627 62.35 -6.45 4.8 j0535+2203 83.75 22.05 20.8 crab nebula j1105+3821 166.25 38.35 14.1 mrk 421 j1654+3945 253.55 39.75 9.4 mrk 501 j1839−0627 279.95 -6.45 6.0 hess j1841-055 j1907+0627 286.95 6.45 5.3 hess j1908+063 j1910+0720 287.65 7.35 4.3 j1912+1026 288.05 10.45 4.2 hess j1912+101 j2021+4038 305.25 40.65 4.3 ver j2019+407 j2031+4157 307.95 41.95 6.1 mgro j2031+41 tev j2032+4130 j1841-0332 280.25 -3.55 4.2 2.1 sky upper limits excluding the sources listed in table 1, we can set upper limits to the γ-ray flux from all other directions in the sky. to estimate the response of the argo-ybj detector we simulated a source located at different declinations, with a power law spectrum in the energy range 10 gev 100 tev and different spectral indices. the number of events is transformed into a flux using the results of the simulation. the 95% c.l. upper limits to the flux of γ-rays with energies above 500 gev for each bin are obtained. the upper limits as a function of the declination are shown in fig. 4 for different photon spectral indices. the limits range between 9% and 44% icrab and are the lowest obtained so far. the lowest limit for a spectral index −2.0 (−3.0) is 5% (9%) icrab, where the crab unit is defined as 5.77×10−11 cm−2 s−1. 3 cosmic ray anisotropy the cr arrival direction distribution and its anisotropy has been a long-standing problem ever since the 1930s. in fact, the study of the anisotropy is a powerful tool to investigate the acceleration and propagation mechanism determining the cr world as we know it. the anisotropy in the cr arrival direction distribution have been observed by different experiments with increasing sensitivity and details at different angular scales. current experimental results show that the main features of the anisotropy are uniform in the energy range (1011 1014 ev), both with respect to amplitude (10−4 10−3) and phase ((0 4) hr). the existence of two distinct broad regions, one showing an excess of crs (called “tail-in”), distributed around 40◦ to 90◦ in r.a., the other a deficit (the “loss cone”), distributed around 150◦ to 240◦ in r.a., has been clearly observed. dec (deg) -10 0 10 20 30 40 50 60 70 9 5 % c .l . u p p e r l im it ( c ra b u n it ) -210 -110 1 =-3.0α =-2.6α =-2.0α figure 4: 95% c.l. flux upper limits for energy above 500 gev, averaged over the right ascension, as a function of the declination. the different curves indicate a different power-law spectral index. the origin of the cr anisotropy is still unknown. unlike predictions from diffusion models, the cr arrival distribution in sidereal time was never found to be purely dipolar. even 2 harmonics were necessary to properly describe the r.a. profiles, showing that the cr intensity has quite a complicated structure unaccountable simply by kinetic models. in the last years the milagro (abdo et al. 2008) and argo-ybj (di sciascio, 2013) collaborations reported evidence of the existence of a medium angular scale anisotropy contained in the tail-in region. the 153 giuseppe di sciascio observation of similar small scale anisotropies has been recently claimed also by the icecube experiment (abbasi et al., 2011) in the southern hemisphere. in fig. 5 the argo-ybj sky map in galactic coordinates as obtained with 4.5 years data is shown. the color scale gives the statistical significance of the observation in s.d. . the map center points towards the galactic anti-center. the maps have been smoothed with an angle given by the psf of the detector for crinduced showers. 0◦360◦ -15 -12 -9 -6 -3 0 3 6 9 12 15 figure 5: argo-ybj sky-map in galactic coordinates. the statistical significance of the observation in s.d. is shown. the map center points towards the galactic anti-center. data have been recorded in 1587 days out of 1656, for a total observation time of 33012 hrs (86.7% dutycycle). a selection of high-quality data reduced the data-set to 1571 days. the zenith angle cut (θ ≤ 50◦) selects the dec. region δ ∼ -20◦÷ 80◦. according to the simulation, the median energy of the isotropic cr proton flux is e50p ≈1.8 tev (mode energy ≈0.7 tev). no gamma/hadron discrimination algorithms have been applied to the data. therefore, the sky map is filled with all crs possibly including photons, without any discrimination. figure 6: one-dimensional projection in right ascension of the two-dimensional cr sky map in local solar time. the red line shows the best-fit to argo-ybj data (crosses). in spite of the fact that the bulk of snr, pulsars and other potential cr sources are in the inner galaxy surrounding the galactic centre, the excess of cr is observed in the opposite, anti-centre direction. as stressed in erlykin & wolfendale (2013), the fact that the observed excesses are in the northern and in the southern galactic hemisphere, favors the conclusion that the cr at tev energies originate in sources whose directions span a large range of galactic latitudes. the right side of the map is full of few-degree excesses not compatible with random fluctuations (the statistical significance is up to 7 s.d.). the observation of these structures is reported by argo-ybj for the first time. so far, no theory of crs in the galaxy exists which is able to explain both large scale and few degrees anisotropies leaving the standard model of crs and that of the local galactic magnetic field unchanged at the same time. 3.1 the compton-getting effect the origin of cr anisotropies is still unknown therefore, the observation of an expected anisotropy is important to check the reconstruction algorithms, the exposure and background calculations and the stability of the detector performance. a well-known expected anisotropy is the so-called compton-getting (cg) effect, a dipole anisotropy in the local solar frame, due to the earth’s motion around the sun (compton & getting, 1935). a significant signal compatible with cg is seen by argo-ybj in solar time above ∼ 8 tev to avoid additional effects due to heliospheric magnetic field and solar activity. in fact, we found that including lower energy events results in much larger modulation amplitudes than those obtained when these events were excluded. fig. 7 shows the solar variations observed by argo-ybj together with the sinusoidal curve best fitted to the data. the fair agreement between data and calculations (φ = 6:00 hr, a = 9.7 · 10−5) make us confident about the capability of argo-ybj in detecting anisotropies at a level of 10−4. 4 prospects: the lhaaso experiment a new experiment has been proposed by the chinese community to face the open problems in galactic cosmic ray physics. the lhaaso experiment is a multicomponent extensive air shower array constituted by: (1) an array consisting of 5137 scintillators (1 m2 each) 15 m away from each other (km2a) and 1200 muon detectors (40 m2 each). the total effective area of the muon detector is about 48,000 m2. (2) a water cherenkov detector array (wcda) consisting of 4 water ponds, 150 × 150 m2 each. the pond depth is 154 argo-ybj: highlights and prospects about 4.5 m. each pond is subdivided into 30 × 30 = 900 cells sized 5 × 5 m2 each, separated by black plastic curtains. an 8 inches pmt looks upward at the bottom of each cell to collect cherenkov photons produced by secondary charged particles in the water pond. (3) 24 wide field of view cherenkov telescope array (wfcta). (4) shower core detector array (scda) with an effective area of 5000 m2. figure 7: layout of the lhaaso experiment. the water ponds allow to improve the sensitivity to γ-ray sources down to a percent of the crab nebula flux in the tev energy region. the km2a array will extend the search for γ-ray sources in the 100 tev region with an unprecedented sensitivity. in fact, exploiting the shower muon content measurement, the detection of photon-induced showers is basically background free above few tens tev. the sensitivity of lhaasowcda + lhaaso-km2a (green line) for detection of point gamma ray sources is compared to other experiments or projects in fig. 8. the observation times is 1 year and 50 hour for wide field-of-view detectors and iact, respectively. the wfcta and the scda will allow, in addition, to study the cosmic ray physics up to the 1018 ev region, thus investigating the transition between galactic and extra-galactic cr components and the elemental composition above pev energies. the different elements of the experiment (scintillators, water pool, wide field of view cerenkov telescopes, neutron detectors) have been successfully tested at the yangbajing laboratory exploiting the cr beam provided by the argo-ybj detector. the installation of the detectors is expected to start between 2-3 years and finish in about 5 years. the proposed site is located in china, in the yunnan province, at an altitude of about 4300 m a.s.l. . figure 8: the sensitivity of lhaaso-wcda + lhaaso-km2a (green line) compared to other experiments or projects. the observation times is 1 year and 50 hour for wide field-of-view detectors and iact, respectively. 5 conclusions the argo-ybj detector exploiting the full coverage approach and the high segmentation of the readout is imaging the front of atmospheric showers with unprecedented resolution and detail. the digital and analog readout will allow a deep study of the cr phenomenology in the wide tev pev energy range. the results obtained in the low energy range (below 100 tev) predict an excellent capability to address a wide range of important issues in astroparticle physics. in this paper we summarized the last results in gamma-ray astronomy and in the study of the cr anisotropy. the new experiment lhaaso, mainly driven by the chinese community, to study the cosmic ray physics up to 1018 ev has been introduced. references [1] abbasi r. et al.: 2011, apj 740, 16. [2] abdo a.a. et al.: 2008, phys. rev. lett. 101, 221101. [3] aielli, g. et al.: 2006, nim a562, 92. [4] aielli g. et al.: 2009, nim a608, 246. doi:10.1016/j.nima.2009.07.020 [5] aielli, g. et al.: 2010a, apj 714, l208 doi:10.1088/2041-8205/714/2/l208 [6] amenomori m. et al.: 2010, astrophys. space sci. trans. 6, 49. doi:10.5194/astra-6-49-2010 [7] bartoli, b. et al.: 2011a, phys. rev. d84, 022003. doi:10.1103/physrevd.84.022003 155 http://dx.doi.org/10.1016/j.nima.2009.07.020 http://dx.doi.org/10.1088/2041-8205/714/2/l208 http://dx.doi.org/10.5194/astra-6-49-2010 http://dx.doi.org/10.1103/physrevd.84.022003 giuseppe di sciascio [8] compton a.h. and getting i.a., 1935, phys. rev. 47, 817. [9] di sciascio g., 2013, epj 52, 04004. [10] erlykin a.d. and wolfendale a.w.: 2013, arxiv:1303.2889. [11] chen, s. et al.: 2013, icrc 2013, id 586. discussion c. munoz-tunon: could you extend a little on the details of the site for the new experiment in china ? g. di sciascio: in principle the new project lhaaso will be located in two different high altitude sites. the lawca (large water cherenkov array) experiment is a possible upgrade of the argo-ybj experiment at the yangbajing laboratory in tibet. the detector will consist in a large water pond l-shaped around the argo-ybj building with an area of about 23,000 m2 and is focused to study gamma-ray astronomy between 100 gev and 30 tev. the detector structure is identical to the wcda of the lhaaso project, therefore this experiment is the phase-0 of the lhaaso project. the data taking is expected to start a couple of years after the start of the construction. j. beall: will the new facility be at the same altitude ? g. di sciascio: yes, the lhaaso experiment will be located at an altitude of about 4300 m a.s.l. in the yunnan province, similar to the altitude of the yangbajing laboratory where is located the argo-ybj detector. c. pittori: can you say something more about the possible connection between observed anisotropies and the heliosphere ? g. di sciascio: as discussed in (amenomori et al., 2010), the main regions of the medium scale anisotropy can be described as two intensity enhancements placed along the hydrogen deflection plane, which contains the directions of the interstellar wind velocity and the interstellar magnetic field surrounding the heliosphere, each symmetrically centered away from the heliotail direction. the separation angle between the heliotail direction and each enhancement monotonously decreases with increasing energy in an energy range 4 30 tev. the msa being placed along the hdp suggests that it is possibly caused by the modulation of galactic cosmic rays in the magnetic field of the heliotail within ∼ 70 au to ∼ 340 au from the sun. 156 the argo-ybj experiment northern sky survey sky upper limits cosmic ray anisotropy the compton-getting effect prospects: the lhaaso experiment conclusions 90 acta polytechnica ctu proceedings 1(1): 90–95, 2014 90 doi: 10.14311/app.2014.01.0090 cosmological evolution of the central engine in high-luminosity, high-accretion rate agn matteo guainazzi1 1european space astronomy centre (esac), p.o. box, 78, e-28691 villanueva de la cañada, madrid, spain corresponding author: matteo.guainazzi@sciops.esa.int abstract in this paper i discuss the status of observational studies aiming at probing the cosmological evolution of the central engine in high-luminosity, high-accretion rate active galactic nuclei (agn). x-ray spectroscopic surveys, supported by extensive multi-wavelength coverage, indicate a remarkable invariance of the accretion disk plus corona system, and of their coupling up to redshifts z ' 6. furthermore, hard x-ray (e ∼>10 kev) surveys show that nearby seyfert galaxies share the same central engine notwithstanding their optical classification. these results suggest that the high-luminosity, high accretion rate quasar phase of agn evolution is homogeneous over cosmological times. keywords: active galactic nuclei. 1 outline thanks to its leap in sensitivity by over two orders of magnitude, the medium sensitivity survey carried out by the einstein observatory (emss, maccacaro et al. 1981) collected for the first a statistically sizable sample of extragalactic x-ray sources. studying a sample of 190 extragalactic sources, giovannelli & polcaro (1986) concluded that the observed linear relation between xray luminosity and redshift over several orders of magnitude made evident a physical continuity between the different classes of extragalactic x-ray objects. the sample in giovannelli & polcaro (1986) was a “mixed bag” of different objects: active galactic nuclei (agn), elliptical/s0 galaxies, and spiral/irregular galaxies. while it is now accepted that most of the x-ray emission in inactive galaxies is due to a combination of diffuse inter-stellar medium, halos, and x-ray binaries (see fabbiano 1989 for an early review), there is almost undisputed consensus that the formidable energy output emerging from agn is due to accretion onto super-massive black holes (see, however, the contribution by prof. kundt for a different view). sub-arcsecond resolution imaging with chandra, as well as the unprecedented xmm-newton throughput have made possible a detailed characterisation of the high-energy spectral energy distribution (sed) in large samples of agn, covering a range in luminosity of almost ten orders of magnitudes, and cosmological distances well beyond the peak of the nuclear activity (brandt & hasinger 2005, elvis et al. 2012, pounds 2013). time is ripe to address some fundamental questions on the nature of accretion onto super-massive black holes, that the small size of the emss extragalactic sample left open. in this contribution i will focus on two of them: • do agn share the same engine at all cosmological times? • do agn share the same engine? the answer for all types of agn at all accretion rates would be clearly: “they don’t”. however, similarities in the observational properties of specific subclasses of agn allow us to give these question astrophysically useful positive answers, at least as far as the members of these sub-classes are concerned. in this review i deal with mainstream radio-quiet1, higheddington ratio (ṁ/ ˙medd ∼> 10−2)2 agn. i assume hereby the traditional wisdom separation in bolometric luminosity between seyfert galaxies (l < 1044 erg s−1) and quasars (qso), whenever relevant. the contribution by m.elitzur to these proceedings discusses a central engine unification scheme encompassing also lowluminosity, low-accretion rate systems. 1in this paper, i assume a qualitative definition of radio-quiet (as opposed to radio-loud) agn as those where the contribution of relativistic jets to the optical-to-x-ray spectral energy distribution is negligible. quantitative definitions of agn radio-loudness were introduced by kellermann et al. (1989), miller et al. (1990), and panessa et al. (2007), among others. 2we define the mass accretion rate in units of eddington as: ṁ ≡ ηledd/c2, where ledd ≡ 4πgmbhmpc, mbh is the black hole mass, mp is the proton mass, c is the speed of light, and η ' 0.1 is the efficiency of gravitational losses into radiation conversion. 90 http://dx.doi.org/10.14311/app.2014.01.0090 cosmological evolution of the central engine in high-luminosity, high-accretion rate agn 2 cosmological evolution of the agn central engine 2.1 direct observables agn are multi-wavelength machines. they emit comparable power per decade frequency from the ir to the γ-ray band (elvis et al. 1994). more importantly for the astrophysical study of accreting super-massive black holes, the bulk of the radiation emitted in each rest-frame energy band carries information on a specific astrophysical process occurring in the nuclear region. uv and x-rays are the most appropriate bands to study spectroscopically the sub-pc scale around accreting massive black holes. in agn accreting at a rate ṁ ∼>10−2 the eddington value the uv emission is dominated by a “blue bump”, believed to be due to thermal emission from the accretion disk (czerny & elvis 1987). the high-energy band is dominated by comptonisation of disk photons by a ∼102 kev temperature corona (zdziarski et al. 1995, perola et al. 2002) with a largely unknown geometry. haardt & maraschi (1991) proposed a “sandwich” geometry embracing the accretion disk. compact coronal geometries have recently experienced a revival following the possible discovery of relativistic light bending (miniutti & fabian 2004) and disk reverberation (fabian et al. 2009) in x-ray observations of nearby seyfert galaxies. uv and x-rays are therefore the natural energy bands where to look for any cosmological evolution of the central engine in luminous agn. in practical terms, metrics accessible to ccd-resolution agn spectra are the intrinsic shape of the x-ray spectrum, parametrised through the photon index γ, which yields information on the corona; and the uv to x-ray flux ratio, parametrised through the αox ∝ log(l2kev /l 2500å ) (tananbaum et al. 1979), which yields information on the disk-corona coupling. for unsaturated comptonisation of non relativistic thermal distributions of electrons, γ ∝ y−1/2, where the comptonisation parameter y ∝ kt × max(τes,τ2es), kt is the electron temperature and τes the optical depth for electron scattering (rybicki & lightman 1979). the measurements of the spectral exponential cut-off at energies e ∼ kt would provide a full characterisation of the physical properties of the corona. however, such measurements are accessible only for a handful of x-ray bright nearby agn even in the nustar (harrison et al. 2013) era. other observables probe in principle the innermost regions of the accretion flow, such as the ubiquitous, and black hole mass dependent, x-ray variability (mc hardy et al. 2006, ponti et al. 2012) and the relativistic distortion of x-ray emission lines produced in a x-ray illuminated accretion disk within a few gravitational radii from the innermost circular stable orbit (fabian et al. 1989). these effects are, however, also difficult to measure at high redshift. fig. 1 shows a qualitative scheme of the intrinsic agn uv-to-x-ray sed for reference. figure 1: schematic view of an agn intrinsic uvto-x-ray spectrum. blue: thermal emission from a luminous accretion disk with a temperature at the innermost radius of 30 ev; red: unsaturated comptonisation of soft disk photons with kte=100 kev and τes=1; magenta: reflection from a plane-parallel infinite slab of cold (∼104k) matter (compton continuum plus fe and nickel kα and kβ fluorescent emission lines; after nandra et al. 2007) sizable samples of x-ray selected agn up to redshift '6 have been collected by difference catalogues and surveys: champ (green et al. 2012), chandra deep fields (steffen et al. 2006), 2xmm (watson et al. 2009, young et al. 2009), cosmos (elvis et al. 2012, lanzuisi et al. 2013). they agree in finding no evidence for cosmological evolution of either γ or αox. similar results were obtained by specific observational programs targeting the farthest x-ray selected qso (vignali et al. 2005). the constrains on αox are particularly tight: once its dependency on the luminosity is corrected for, the average of the αox distributions in different redshift bins agree within ±0.1 (steffen et al. 2006). taken as a whole, these results indicate a remarkable invariance of the corona-disk coupling over 90% of the look-back time. 2.2 indirect observables other observables can be used to probe indirectly the nuclear sed: we focus in this section on emission lines in the optical, and in the x-ray band. optical spectroscopy probes the physical condition of gas on a variety of different scales in the nuclear environment: from the broad line regions (blrs) clouds, emitting lines primarily from permitted transitions with widths ∼>2000 km s−1, to the narrow line regions 91 matteo guainazzi (nlrs), emitting lines primarily from forbidden transitions with widths ∼<1000 km s−1. these gaseous systems cover the whole range of distances from the central engine between light-days (peterson et al. 2004) to tens-hundreds of kpc (pogge 1988, 1989a, 1989b). extended nlrs are also copious sources of x-rays (young et al. 2001, bianchi et al. 2006). gas in these regions is photo-ionised by the agn. photo-ionisation models applied to the observed spectra allows one to reconstruct the ionising sed, and even in a few cases the past history of the accreting black hole (dadina et al. 2010). the farthest optically identified agn is ulasj1120+0641 at z = 7.085, a mbh = 2 × 109m� quasar (qso). its optical spectrum is almost indistinguishable from composite spectra of agn at lower redshift (mortock et al. 2011). studies of qso optical spectra at z ≥ 5 indicate no evolution of the metallicity as a function of redshift (juarez et al. 2009). these results suggest that the invariance of the central engine extends to the immediate environs of the accreting black hole up to redshifts '7. due to the combination of high fluorescent yield and abundance, iron kα fluorescent emission line is among the most common spectral features in x-ray spectra of cosmic sources. it is well isolated from transitions due to the closest elements; even at ccd resolution (e/∆e∼40) one can easily resolve fluorescence emission lines from he-like, and h-like recombination lines (statistics permitting). an unresolved component of the neutral iron kα fluorescent line has been detected almost ubiquitously in nearby agn (yaqoob & padmanabhan 2004; nandra 2006). this feature traces reprocessing of the primary nuclear emission by optically thick gas. while this feature is too weak to be detected in x-ray spectra of individual sources at high redshift, spectra stacking techniques allowed the cosmos survey to probe its presence up to redshift z ' 3. fitting these stacked spectra with simple phenomenological continua leaves residuals consistent with neutral iron kα emission (and weaker residuals consistent with recombination lines from heand h-like iron). the equivalent width of such features is independent of redshift (chaudhary et al., 2010; iwasawa et al. 2012), once corrected for the known dependence on the absorption-corrected x-ray luminosity, the so called “iwasawa-taniguchi” effect (iwasawa & taniguchi 1993; bianchi et al. 2007). this result implies a similarity of the primary x-ray sed, as well as of the geometrical configuration and ionisation stage of the reprocessing matter at different cosmological times. 2.3 however, agn do evolve ... these results shall not be interpreted as due to the fact that the agn population as a whole do not evolve! the agn population evolves in density and luminosity over cosmological times (see brandt & hasinger 2005 for a recent review). this may in principle affect the agn population being probed at different redshift by current surveys. hopkins et al. (2008) and hickox et al. (2009) present an evolutionary scenario whereby all agn go through the same evolutionary steps. merging or secular processes (gas instabilities) trigger simultaneously the quasar phase and the growth of the stellar bulge when a galaxy’s dark matter halo reaches a critical mass ∼1012−13 m�. the peak of the quasar phase accreting at high eddington ratio, when massive black holes accrete the bulk of their mass, occurs at different redshifts depending on the mass of the primeval dark mass halo. the whole agn evolutionary sequence would start from a cold gas-rich, rotation dominated galaxy, pass through a dustand optically-thick gas enshrouded phase, evolve into the qso phase, and then to a gradual decline of the agn activity until the galaxy become a “red and dead” elliptical with intermittent radio activity. all these phases would be present in an observed agn sample at any redshift. the qso phase is, however, the best one to probe the conditions of the innermost regions close to the central engine. the discovery of a ∼109m� black hole at z ∼ 7 is puzzling for models of cosmological black hole growth. it is not simple to understand how the required mass could have accreted from seed black holes at z ≥ 10. broadly speaking (see volonteri 2010 for a review) models of seed black hole formation and evolution can be classified as: a) light (mbh ∼102−3 m�) and rapidly accreting seeds forming at very early cosmic time (z '20–50); b) heavy (mbh ∼104−6 m�) and slowly accreting seeds forming at later stages (z '5– 10): c) intermediate seeds between the two cases above. none of these scenarios is free of potential shortcomings (milosavljević et al. 2009; volonteri & begelman 2010). if massive enough (∼>105 m�), electromagnetic signals from these early seeds could be detected by future large observatories. otherwise the detection of gravitational waves produced when compact objects fall onto the seed could provide constraints on a population in the mass range mbh ∼104−7 m�. these topics will be addressed by a future esa’s l-class x-ray mission in the framework of the cosmic vision program, one of whose basic science theme is: “how did the universe originate and what is it made of?” (esa 2005). 3 do really agn share the same engine? so far we have assumed the all agn in the quasar phase at a given time share the same central engine. x-rays provide the ideal wavelength to test this assumption for radio-quiet agn in the local universe. in order to 92 cosmological evolution of the central engine in high-luminosity, high-accretion rate agn understand why, we need to have a closer look at the configuration of gas and dust surrounding the agn central engine, and on how it affects the phenomenological appearance of agn. in sect. 2.2 it was shown that the ubiquitous detection of kα fluorescent emission lines corresponding to neutral or mildly ionised iron (together with the associated continuum compton-reflection “hump”; nandra & pounds 1994) is considered a clear evidence for reprocessing of the primary nuclear emission by optically thick gas in the agn environs. the geometry of this gas structure is controversial. traditional wisdom invoked a compact symmetrical molecular “torus” (antonucci & miller 1985). its azimuthal symmetry would introduce a dependence of the agn observational properties on the line-of-sight to the nucleus. these orientation effects would be primarily (but not exclusively; see nicastro 2000) responsible for the properties of their optical spectra: blrs would be visible only through lines-ofsight not intercepting the torus. for the high reddening expected when radiation passes through the torus clouds (nh > 10 22 cm−2) optical broad lines would be suppressed, and only narrow emission lines would be visible (see antonucci 1993 for a review). in this “agn unified scenario” broadand narrow-line agn share the same common engine. hoverer, early simulations (pier & krolik 1992) pointed out that it is difficult to prevent such a structure from gravitationally collapsing on time-scales much shorter than the typical agn duty-cycle (∼>1 gyr). recent models invoke a ”clumpy” model as a solution to this issue (nenkova et al. 2002, 2008). in this framework, the agn classification into broadand narrow-line agn should be reinterpreted in probabilistic terms: even agn seen at very high inclination angles (i.e., for which the traditional unified scenario would predict that they could be only narrow-line agn), would have some probability of being detected as broad-line agn depending on the number, size and dynamics of the torus clouds (eliztur 2012; see also eliztur’s contribution to these proceedings for a discussion of this interpretative scenario). first direct x-ray evidence for the torus clumpiness might have been recently detected in the heavily obscured agn markarian 3 (guainazzi et al. 2012). notwithstanding the detailed structure of the torus, and the consequently corrected interpretation of the unified scenario, there is a tight correlation between optical and x-ray spectroscopic properties in nearby agn: optical “narrow-line” agn exhibit x-ray spectra seen through large column densities of cold photoelectrically absorbing gas (nh ∼> 1022 cm−2; awaki et al. 1991; risaliti 2002); optical “broad-line” agn are x-ray unobscured. exceptions to this rules are '3% in large agn x-ray spectroscopic surveys (mateos et al. 2010), and can be easily explained by variability between non-simultaneous optical and x-ray observations, x-ray obscuration by matter in the host galaxy, or even classification issues in poor-quality optical spectra. one needs to carefully correct for absorption (singh et al. 2011), or observe in energy ranges (as much as possible) unaffected by obscuration to probe the innermost region of agn belonging to different optical classes. the latter avenue have been recently opened by the unprecedented sensitivity achieved by the integral/ibis and the swift/bat instruments above 10 kev. sample of more than 100 nearby agn detected above 10 kev are discussed, among others, by beckmann et al. (2009); ricci et al. (2011), molina et al. (2013) for ibis, and burlon et al. (2012) for bat, respectively. the distributions of γ in samples of x-ray obscured and unobscured agn are consistent within ∆γ '±0.1. the distributions of αox remain indistinguishable also when the x-ray luminosity density is calculated in the hard x-ray band (beckman et al. 2009). interestingly enough, ricci et al. (2011) suggested that the reprocessing continuum spectra components are stronger when agn are obscured by a column density 1023 cm−2≤nh≤1024 cm−2. this ingredient does not belong to the 0th order unified scenarios. at this level of obscuration a contribution by blr clouds is likely (elvis et al. 2004; risaliti et al. 2005, 2007; bianchi et al. 2009). while this result is a wise reminder of the intrinsic complexity of gas and dust structures responsible for orientation(and time-)dependent obscuration of the primary continuum, it does not in itself invalidate the basic assumptions of the unified scenario. 4 conclusions in summary: • hard x-ray spectroscopy shows that all (local) high-accretion rate agn share the same central engine notwithstanding the optical classification • uv and x-ray spectroscopy shows that agn in the same phase of their evolution (the high accretion rate, unobscured “quasar phase”) share the same engine up to z ' 6 • optical (blr emission) and x-ray spectroscopy 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done c., smith d., et al., 1995, apj, 438, 63 95 outline cosmological evolution of the agn central engine direct observables indirect observables however, agn do evolve ... do really agn share the same engine? conclusions acta polytechnica ctu proceedings doi:10.14311/app.2015.1.0001 acta polytechnica ctu proceedings 2:1–7, 2015 © czech technical university in prague, 2015 available online at http://ojs.cvut.cz/ojs/index.php/app reinforced encoding for planning as sat tomáš balyo∗, roman barták, otakar trunda department of theoretical computer science and mathematical logic, faculty of mathematics and physics, charles university, malostranske namesti 2/25, praha 1, czech republic ∗ corresponding author: biotomas@gmail.com abstract. solving planning problems via translation to satisfiability (sat) is one of the most successful approaches to automated planning. we propose a new encoding scheme, called reinforced encoding, which encodes a planning problem represented in the sas+ formalism into sat. the reinforced encoding is a combination of the transition-based sase encoding with the classical propositional encoding. in our experiments we compare our new encoding to other known sas+ based encodings. the results indicate, that he reinforced encoding performs well on the benchmark problems of the 2011 international planning competition and can outperform all the other known encodings for several domains. keywords: planning, satisfiability, encoding. 1. introduction planning is the problem of finding a sequence of actions – a plan, that transforms the world from an initial state to a state that satisfies some goal conditions. the world is fully-observable, deterministic and static (only the agent we make the plan for changes the world). the number of possible states of the world as well as the number of possible actions is finite, though possibly very large. we will assume that the actions are instantaneous (take a constant time) and therefore we only need to deal with their sequencing. actions have preconditions, which specify in which states of the world they can be applied, as well as effects, which dictate how the world will be changed after the action is executed. one of the most successful approaches to planning is encoding the planning problem into a series of satisfiability (sat) formulas and then using a sat solver to solve them. the method was first introduced by kautz and selman [1] and is still very popular and competitive. this is partly due to the power of sat solvers, which are getting more efficient year by year. since then many new improvements have been made to the method, such as new compact and efficient encodings [2–5], better ways of scheduling the sat solvers [3] or modifying the sat solver’s heuristics to be more suitable for solving planning problems [6]. in this paper we present a new encoding scheme. it is inspired by the sase transition-based encoding [2], which was the first sat encoding based on the sas+ planning formalism. the motivation for our work is to make the sase encoding more robust by incorporating the strengths of older encoding schemes. we will prove the correctness of our encoding and compute an upper bound on the size of the encoded formula. in the experimental section of the paper we compare our new encoding to other sas+ encodings on benchmark problems from the 2011 international planning competition (ipc) [7]. 2. preliminaries in this section we give the basic definitions of satisfiability, and planning with parallel plans. 2.1. satisfiability a boolean variable is a variable with two possible values true and false. a literal of a boolean variable x is either x or ¬x (positive or negative literal). a clause is a disjunction (or) of literals. a clause with only one literal is called a unit clause and with two literals a binary clause. an implication of the form x ⇒ (y1 ∨ ·· · ∨ yk) is equivalent to the clause (¬x∨y1∨·· ·∨yk). a conjunctive normal form ( cnf) formula is a conjunction (and) of clauses. a truth assignment φ of a formula f assigns a truth value to its variables. the assignment φ satisfies a positive (negative) literal if it assigns the value true (false) to its variable and φ satisfies a clause if it satisfies any of its literals. finally, φ satisfies a cnf formula if it satisfies all of its clauses. a formula f is said to be satisfiable if there is a truth assignment φ that satisfies f. such an assignment is called a satisfying assignment. the satisfiability problem (sat) is to find a satisfying assignment of a given cnf formula or determine that it is unsatisfiable. 2.2. planning in the introduction we briefly described what planning is, in this section we give the formal definitions. we will use the multivalued sas+ formalism [8] instead of the classical strips formalism [9] based on propositional logic. a planning task π in the sas+ formalism is defined as a tuple π = {x,o,si,sg} where • x = {x1, . . . ,xn} is a set of multivalued variables with finite domains dom(xi). 1 http://dx.doi.org/10.14311/app.2015.1.0001 http://ojs.cvut.cz/ojs/index.php/app t. balyo, r. barták, o. trunda acta polytechnica ctu proceedings • o is a set of actions (or operators). each action a ∈ o is a tuple (pre(a), eff(a)) where pre(a) is the set of preconditions of a and eff(a) is the set of effects of a. both preconditions and effects are of the form xi = v where v ∈ dom(xi). • a state is a set of assignments to the state variables. each state variable has exactly one value assigned from its respective domain. we denote by s the set of all states. si ∈ s is the initial state. sg is a partial assignment of the state variables (not all variables have assigned values) and a state s ∈ s is a goal state if sg ⊆ s. an action a is applicable in the given state s if pre(a) ⊆ s. by s′ = apply(a,s) we denote the state after executing the action a in the state s, where a is applicable in s. all the assignments in s′ are the same as in s except for the assignments in eff(a) which replace the corresponding (same variable) assignments in s. if p = [a1 . . .ak] is a sequence of actions, then apply(p,s) = apply(ak, apply(ak−1 . . . apply(a2, apply(a1,s)) . . . )). a sequential plan p of length k for a given planning task π is a sequence of k actions p such that sg ⊆ apply(p,si). 2.3. parallel plans a parallel plan p with makespan k for a given planning task π is a sequence of sets of actions (called parallel steps) p = [a1, . . . ,ak] such that e(a1)⊕···⊕e(ak) is a sequential plan for π, where e is an ordering function, which transforms a set of actions ai into a sequence of actions e(ai) and ⊕ denotes the concatenation of sequences. let us denote by sj the world state in between the parallel steps aj and aj+1, which is obtained by applying the sequence e(aj) on sj−1, i.e., sj = apply(e(aj),sj−1) (except for s0 = si). in this paper we will use the ∀-step parallel planning semantics [10], which requires that each action a ∈ aj is applicable in the state sj, the effects of all actions are applied in sj+1 and all possible orderings e of the sets aj make valid sequential plans (hence the name ∀-step semantics). to ensure, that each ordering of the sets of actions in a parallel plan leads to a valid sequential plan, it is sufficient to check that the actions in each set are pairwise independent [3]. we say that two actions a1 and a2 are independent if they do not share common variables, i.e., scope(a1) ∩ scope(a2) = ∅, where scope(a) ⊆ x is a set of all state variables that appear in pre(a) and eff(a). note, that the pairwise independence of actions is a sufficient but not a necessary condition for the parallel steps in a ∀-step semantics plan, as the following example demonstrates. example 1. let a1 and a2 be two actions such that pre(a1) = pre(a2) = {x = 1}, eff(a1) = {y = 2}, and eff(a2) = {z = 2}. clearly, a1 and a2 are not sp1 planningassat (π) sp2 k := 0 sp3 repeat sp4 k := k + 1 sp5 f := encodetaskwithmakespan(π, k) sp6 until issatisfiable(f) sp7 p := extractplan(getsatassignment(f)) sp8 return p figure 1. pseudo-code of the basic planning as satisfiability algorithm. independent (they share the variable x), however, they can be ordered arbitrarily to achieve the same changes between two given states. the pairwise independence of actions in each step of a parallel plan also implies that they can be executed in parallel (at the same time). 3. finding plans using sat the basic idea of solving planning as sat is the following [1]. we construct (by encoding the planning task) a series of sat formulas f1,f2, . . . such that fi is satisfiable if there is a parallel plan of makespan ≤ i. then we solve them one by one starting from f1 until we reach the first satisfiable formula fk. from the satisfying assignment of fk we can extract a plan of makespan k. the pseudo-code of this algorithm is presented in figure 1 the method was first introduced by kautz and selman [1] and is still very popular and competitive. this is partly due to the power of sat solvers, which are getting more efficient year by year. since then many new improvements have been made to the method, such as new compact and efficient encodings [2–5], better ways of scheduling the sat solvers [3] or modifying the sat solver’s heuristics to be more suitable for solving planning problems [6]. clever ways of solver scheduling [3] can significantly improve the performance of the planning algorithm at the cost of possibly longer-makespan plans. nevertheless, we will use the basic one-by-one scheduling since we are interested only in comparing the properties of encodings, i.e., the construction of the formulas. in the following section we will describe how a formula encoding a planning task can be constructed using our new reinforced encoding. 4. reinforced encoding our goal is (given a planning task π = {x,o,si,sg} and an integer k) to construct a cnf formula fk such that fk is satisfiable only if there is a parallel plan of at most k steps for π. we also want to construct fk in a way, that in the case it is satisfiable, we can easily extract a plan from its satisfying assignment. before we describe the formula, we need to introduce the notion of transitions [2]. 2 vol. 2/2015 reinforced encoding for planning as sat a transition represents a change of a state variable x ∈ x from one value to another from its domain dom(x) or from an arbitrary value to a specific value. there are the following three kinds of transitions. • an active transition changes the value of the variable x from d to e such that d 6= e, {d,e}⊆ dom(x), it is denoted by δx: d→e. an action a has an active transition δx: d→e if (x = d) ∈ pre(a) and (x = e) ∈ eff(a). • a prevailing transition conserves the value of the variable x (if it was d, then it remains d, d ∈ dom(x)), it is denoted by δx: d→d. an action a has a prevailing transition δx: d→d if (x = d) ∈ pre(a) and there is no assignment related to x in eff(a). • a mechanical transition changes the value of the variable x from any value to the value d (d ∈ dom(x)), it is denoted by δx: ∗→d. an action a has a mechanical transition δx: ∗→d if (x = d) ∈ eff(a) and there is no assignment related to x in pre(a). example 2. the action a with preconditions pre(a) = {x = 1,y = 3} and effects eff(a) = {y = 1,z = 2} has one active transition (δy: 3→1), one prevailing transition (δx: 1→1), and one mechanical transition (δz: ∗→2). the transition set of an action a is the set of all transitions that a has, it is denoted by ∆a. by ∆p we will mean the set of all possible prevailing transitions of a planning task, i.e., ∆p = {δx: d→d | x ∈ x,d ∈ dom(x)}. the set of all transitions ∆ is the union of all the prevailing transitions and the transition sets of all the actions ∆ = ∆p ∪{∆a | a ∈ o}. by ∆x ⊆ ∆ where x ∈ x we will denote the set of all transitions related to the variable x. the constructed formula fk will have the following three kinds of boolean variables. • action variables ati indicating whether the i-th action is used in the t-th step. we will have one such variable for each action from the description of the planning task and for each of the k parallel steps. • assignment variables btx=v indicating whether the value of the variable x is equal to v in the end of the t-th step (after applying the actions of the t-th step). we will have one such boolean variable for each state variable x ∈ x and each value v ∈ dom(x) for each of the k parallel steps. • transition variables ctδ (or c t x: d→e where δ = δx: d→e) indicating whether the transition δ occurred during the t-th step. we will have one such variable for each δ ∈ ∆ for each of the k parallel steps. now we are ready to define the clauses contained in fk. the following set of binary clauses will enforce, that at most one value is assigned to each state variable x ∈ x. (¬btx=vi ∨¬b t x=vj ) ∀x ∈ x, vi 6= vj, {vi,vj}⊆ dom(x), ∀t ∈{1, . . . ,k} (1) the following three kinds of clauses connect the assignment variables with the transition variables. the first set of clauses ensures that each transition δx: d→e (including prevailing transitions δx: e→e and mechanical transitions δx: ∗→e) implies that x = e at the end of each step. (¬ctδx: d→e ∨ b t x=e) ∀δx: d→e ∈ ∆,∀t ∈{1, . . . ,k} (2) similarly, we need to add clauses for each transition δx: d→e (except for mechanical transitions) to enforce that x = d holds at the end of the previous step, except for the first step, where we explicitly disable all the transitions that are not compatible with the initial state (using the clauses from equation 8). (¬ctδx: d→e ∨ b t−1 x=d) ∀δx: d→e ∈ ∆,d 6= ∗,∀t ∈{2, . . . ,k} (3) the third kind of clauses is needed to guarantee, that if a variable x has the value v then there is a transition which changes the value of x to v. (¬btx=v ∨ c t δ1 ∨·· ·∨ ctδm ) ∀x ∈ x,v ∈ dom(x),δ1, . . . ,δm transform x to v, ∀t ∈{1, . . . ,k} (4) next we describe the clauses that connect the action variables with the transition variables. if an action a is selected, then all the transitions in its transition set ∆a must be selected as well. this implication is expressed via the following clauses. (¬at ∨ ctδ) ∀a ∈ o,∀δ ∈ ∆a,∀t ∈{1, . . . ,k} (5) also we need to make sure, that transitions (except for prevailing transitions) cannot happen without actions that have them in their transition sets. the following set of clauses will ensure this. (¬ctδ ∨a t s1 ∨·· ·∨atsm ) ∀δ ∈ (∆ \ ∆p), support(δ) = {s1, . . . ,sm}, ∀t ∈{1, . . . ,k} (6) by support(δ) we mean the set of indices of actions that have δ in their transition set, i.e., support(δ) = {i | ai ∈ o; δ ∈ ∆ai}. next we need to deal with the interfering actions inside a parallel step. as discussed earlier, it is sufficient to ensure, that only pair-wise independent actions are together in each parallel step. we will achieve this by disabling all pairs of non-independent (interfering) actions. to extrude interfering actions from the parallel 3 t. balyo, r. barták, o. trunda acta polytechnica ctu proceedings steps we will add binary clauses for all the interfering action pairs. (¬ati ∨¬a t j) ∀ai,aj ∈ o, ai,aj not independent, ∀t ∈{1, . . . ,k} there might be a plenty of interfering action pairs producing a lot of clauses. but if we look carefully at the clauses we have already described, we can see, that most of the interfering actions cannot occur together anyway as we will show via the following notion of compatible actions. two sets of conditions (assignments) are compatible if they assign the same values to the variables they share. two actions a1 and a2 are compatible if the preconditions of a1 are compatible with the preconditions of a2 and also the effects of a1 are compatible with the effects of a2. due to the clauses that enforce, that actions imply their transitions (5) and their connection to assignment variables (2 and 3) together with the clauses that forbid a state variable to have more than one value 1, actions that are not compatible cannot be in a parallel step together. therefore it is enough to suppress compatible interfering action pairs. (¬ati ∨¬a t j) ∀ai,aj ∈ o, ai,aj compatible and not independent, ∀t ∈{1, . . . ,k} (7) lastly, we add the clauses that enforce the initial state to hold in the beginning and the goal conditions to be satisfied in the end. as for the initial state, we will disable all the transitions that are not compatible with the initial state, i.e., if a variable x has the value d in the initial state, then all the transitions that change x from a value other than d are disabled by using a unit clause. note, that mechanical transitions are always compatible with the initial state (or any other state) and therefore no mechanical transition is disabled. (¬c1δx: d→e ) ∀δx: d→e ∈ ∆, (x = d) /∈ si (8) to encode the goal conditions we will use unit clauses with assignment variables. fore each goal condition (x = v) ∈ sg we will have a unit clause (bkx=v) which forces the value of x to be v after the last parallel step. (bkx=v) ∀(x = v) ∈ sg (9) the formula fk for the reinforced encoding is a conjunction of the clauses defined in equations 1, 2, 3, 4, 5, 6, 7, 8, and 9. a ∀-step parallel plan can be extracted from any satisfying assignment of fk in the following way. let φ be a satisfying assignment of fk. pφ is a sequence of action sets such that its t-th set contains those actions ai ∈ o for which φ(ati) = true. 4.1. correctness in this subsection the prove the correctness of our encoding, i.e., the following proposition. proposition 1. if the formula fk obtained using the reinforced encoding of the planning task π is satisfied by a truth assignment φ then pφ is a valid ∀-step parallel plan of makespan k for the planning task π. proof. the requirements for the action sets given by the ∀-step semantics are clearly satisfied: • the preconditions of actions in each parallel step are satisfied due to 2, 3, and 5 • the effects are propagated also due to 2, 3, and 5 • the actions can be ordered arbitrarily thanks to 7 it remains to prove that psφ = [e(a1)⊕···⊕e(ak)] is a valid (sequential) plan for π, where ⊕ denotes the concatenation of sequences and e is an arbitrary ordering of an action set. let us observe, that the transitions of the state variables are consistent at each step, i.e., exactly one transition is allowed for each state variable (due to 2, 3, and 1) and a non-prevailing transition cannot happen without an action that has it (thanks to 6). prevailing transitions do not have to be supported by any actions since they are used to preserve the values of the variables that are not changed in the given step by any actions. furthermore, all the transitions between two neighboring parallel steps must be compatible due to 2, 3, 4 and 1. note, that it may happen, that a variable x has no value assigned in the end of a step t (all the btx=v,v ∈ dom(x) are false) and no transition related to the variable is selected (all the ctδ,δ ∈ ∆x are false). however, this can only occur for variables that are not used in the goal conditions or by actions that appear in the t-th step or later. since the action variables imply the proper transition variables thanks to 5, the actions must be applicable if their action variable is true and also the transition connected to the action must happen. thanks to 8 only transitions compatible with the initial state can happen in the first step and because of 2 and 9 only transitions that change the variables to their goal values are allowed in the last step. this fact together with the consistency of the transitions during all the k steps implies the validity of psφ for the planning task π. the reversed implication, which is that if pφ is a valid ∀-step parallel plan of makespan k, then φ satisfies fk, does not hold. this is because pφ may contain a non-independent pair of actions in one of its steps and still be a valid ∀-step plan (see example 1). such a pair of actions would make one of the clauses of type 7 unsatisfied. 4 vol. 2/2015 reinforced encoding for planning as sat 4.2. size of the encoded formula the size of the formulas will of course depend on the parameters of the planning task being encoded. we will use the following quantitative properties of a planning task π = (x,o,si,sg) to compute the upper bounds. • n the number of actions (n = |o|). • v the number of state variables (v = |x|). • d the maximum domain size of any state variable (d = maxx∈x{|dom(x)|}). • p the maximum number of preconditions or effects an action has (p = maxa∈o{|pre(a)|, |eff(a)|}) typically, the number of actions is much higher than the other parameters. from these values we can compute the following upper bounds related to the planning task. • the number of assignments is at most vd. • the number of transitions is at most v(d2 +d) since there are at most d(d− 1) active, d prevailing, and d mechanical transitions for each variable. from these bounds it is apparent, that fk (a formula for makespan k) has at most kn action variables, kvd assignment variables, and kv(d2 + d) transition variables. therefore the total number of boolean variables in fk is k(n + vd(d + 2)). now let us compute an upper bound on the number of clauses in fk. we will count separately the number of unit clauses (clauses with one literal), binary clauses (clauses with two literals), and horn clauses (clauses with at most one positive literal). the formula fk obtained by the reinforced encoding is the conjunction of the clauses defined in equations 1, 2, 3, 4, 5, 6, 7, 8, and 9 • there are at most kvd2 clauses of the type 1 – one for each step and variable and two different values from its domain. these clauses are binary and horn. • there are at most kv(d2 + d) clauses of the both type 2 and type 3 – one for each step and transition. these clauses are binary and horn. • there are at most kvd clauses of the type 4 – one for each step and assignment. • there are at most 2knp clauses of the type 5 – one for each step, action and each of its transitions (there are at most 2p transitions connected to each action). these clauses are binary and horn. • there are at most kv(d2 + d) clauses of the type 6 – one for each step and transition. • there are at most kn2 clauses of the type 7 – one for each step, and each pair of compatible interfering actions (at most each pair of actions). these clauses are binary and horn. • there are at most v(d2 + d) clauses of type 8 – one for each transition that is not compatible with the initial state (at most all the transitions). these are unit clauses. • there are at most v clauses of the type 9 – one for each goal condition. these clauses are unit. in total we have k(n2 +2np+4vd2 +4vd)+vd2 +vd+v clauses, from which vd2 + vd + v are unit clauses and k(n2 + 2np + 3vd2 + 2vd) are both binary and horn clauses. 5. experimental evaluation to evaluate the performance of our new reinforced encoding, we compared it with three other sas+ based encodings of planning as sat. we ran experiments with a 30 minutes time limit using the following four encodings. • reinforced encoding (reinf). a java implementation of our new reinforced encoding as described in the previous section. • direct encoding (dir). we implemented a simple encoding based on the historically first encoding of planning as sat [1]. we adapted it for the sas+ formalism. this encoding is similar to our reinforced encoding but uses only action and assignment variables. • sase encoding (sase). our java implementation1 of the transition-based sase encoding [2]. this encoding uses only action and transition variables. • r2∃-step encoding (r2∃). the original java implementation of the r2∃-step encoding [5]. this encoding differs from the previous three encoding significantly since it uses a different parallel planning semantics. the r2∃-step encoding allows more actions inside the parallel steps, therefore it often finds plans with much lower makespans. lower makespan indicates that fewer sat solver calls are required to find a plan, however, it does not say anything about its length, i.e., the total number of actions it contains. 5.1. experimental setting to compare the performance of the encodings we created a simple script, which iteratively constructed and solved the formulas for time steps 1, 2, . . . until a satisfiable formula was reached (see figure 1). for each encoding we used the same sat solver – lingeling[11] (version ats). the time limit was 30 minutes for the sat solving part, i.e., the total time the sat solver could spend solving the formulas f1,f2, . . . for each problem instance was 30 minutes. the time required for the generation of f1,f2, . . . is usually negligible compared to the time required to solve them and therefore we will ignore it. hence the overall planning time could exceed the given time limit for a problem instance. 1the original sase implementation cannot be used since it does not support the format of the latest benchmark problems 5 t. balyo, r. barták, o. trunda acta polytechnica ctu proceedings domain dir sase reinf r2∃ barman 4 4 4 8 elevators 20 20 20 20 floortile 16 11 18 18 nomystery 20 10 20 6 openstacks 0 0 0 15 parcprinter 20 20 20 20 parking 0 0 0 0 pegsol 10 6 10 19 scanalyzer 14 12 15 9 sokoban 2 2 2 2 tidybot 2 2 2 2 transport 16 17 18 13 visitall 12 9 10 20 woodworking 20 20 20 20 total 156 133 159 172 table 1. the number of problems (out of 20) in each domain that the encodings solved within the time limit (30 minutes for sat solving). the experiments were run on a computer with intel i7 920 cpu @ 2.67 ghz processor and 6 gb of memory. the benchmark problems of the ipc are organized into domains. each domain contains 20 problems and there are 14 domains which results in a total of 280 problems. the benchmark problems are provided in the pddl format, however, the encodings require input in the sas+ format. we used helmert’s translation tool, which is a part of the fast downward planning system [12], to obtain the sas+ files from the pddl files. the translation is very fast requiring only a few seconds for all domains. 5.2. experimental results the number of solved instances in presented in table 1. looking at the results from the perspective of the domains, we can observe, that the elevators, parcprinter, and woodworking domains are entirely solved by every encoding. on the other hand, the parking domain is so difficult that not even a single problem is solved by any of the encodings. the openstacks domain is very difficult for all but the r2∃-step encoding. the sokoban and tidybot domains are also very hard for all of the encodings, only two of the twenty problems are solved by each encoding. if we compare the encodings, we can observe that the r2∃-step encoding has the highest total number of solved instances followed by our new reinforced encoding. as for the individual domains, the r2∃-step encoding solves strictly more problems than the other encodings in four cases. the reinforced encoding achieves this for three domains, while the direct and sase encoding cannot outperform the other encoddomain dir sase reinf r2∃ barman 121 121 121 84 elevators 190 190 190 85 floortile 302 181 344 169 nomystery 347 119 347 30 openstacks 93 parcprinter 261 261 261 30 parking pegsol 222 131 222 158 scanalyzer 83 61 95 17 sokoban 60 60 60 27 tidybot 14 15 15 6 transport 221 242 262 55 visitall 223 110 146 34 woodworking 68 68 68 33 table 2. the sum of makespans of the plans found within the time limit for each domain. lower makespan means fewer sat solver calls, it does not indicate better plan quality. domains with the same number of solved problems for each encoding are highlighted. ings in any of the domains. the reinforced encoding solves the same number of problems as any other encoding in seven cases. except for the visitall domain, the reinforced encoding is never worse than the direct or sase encoding. looking at the makespans of found plans displayed in table 2 we can observe that the makespans for the r2∃-step plans are indeed significantly lower than the makespans of plans found by the other three encodings. as expected, in the cases when the direct, sase, and reinforced encodings solve all the problems (or solve the same problems) their total makespans are identical. this is due to the fact that these three encodings use the same ∀-step parallel planning semantics. the times required to solve the problems are presented in table 3. if we look at the results for the domains, where each encoding solved the same number of problems, i.e., the highlighted domains, we can notice, that except for the parcprinter and sokoban problems, the runtime of the r2∃-step encoding is much higher than the runtime of the other methods. if we also look at table 2, which contains the total makespan of the found plans, we can deduce, that lower makespan, i.e., fewer sat calls does not necessarily mean faster planning, especially not in the case of the easy domains. nevertheless, for the domains, where r2∃-step significantly outperformed the other methods – openstacks, pegsol, and visitall, the makespans are much lower than the makespans of the other methods, despite the fact, that they solved fewer problems. 6 vol. 2/2015 reinforced encoding for planning as sat domain dir sase reinf r2∃ barman 2041.44 1549.31 1680.53 2999.30 elevators 33.96 55.72 38.10 288.29 floortile 7327.15 2083.57 1206.18 1380.63 nomystery 3798.45 1377.23 1894.65 1927.40 openstacks 2679.68 parcprinter 10.07 28.25 12.15 4.24 parking pegsol 3796.11 2096.72 4971.97 830.72 scanalyzer 1401.21 831.80 1435.19 1118.89 sokoban 592.03 1337.87 857.04 550.18 tidybot 75.68 74.02 118.09 480.85 transport 1404.23 3554.90 3203.62 7418.01 visitall 2380.76 683.25 726.53 14.24 woodworking 2.57 2.04 3.84 138.61 table 3. the time in seconds required to solve all the problems that were solved within the time limit. the presented time is the sum of times the sat solver alone required, formula generation time is not included. domains with the same number of solved problems for each encoding are highlighted. 6. conclusion in this paper we have introduced a new encoding of a planning problem represented in the sas+ formalism into sat. our new encoding performs well on the benchmark problems of the 2011 international planning competition. it can strictly outperform all the other evaluated sas+ encodings in three domains and solve the same number of problems as any other encoding for seven domains out of fourteen. on the remaining four domains our encoding is outperformed by the r2∃-step encoding which uses a different parallel planning semantics. as for future work, we believe that the reinforced encoding can be improved by decreasing the number of its clauses by using a more compact way of encoding of the action interference constraints. acknowledgements the research is supported by the czech science foundation under the contract p103/10/1287 and by the grant agency of charles university under contracts no. 600112 and no. 390214. this research was also supported by the svv project number 260 104. references [1] h. a. kautz, b. selman. planning as satisfiability. in ecai, pp. 359–363. 1992. http://citeseerx.ist.psu. edu/viewdoc/summary?doi=10.1.1.35.9443. [2] r. huang, y. chen, w. zhang. a novel transition based encoding scheme for planning as satisfiability. in m. fox, d. poole (eds.), aaai. aaai press, 2010. http://citeseerx.ist.psu.edu/viewdoc/summary? doi=10.1.1.182.9561. [3] j. rintanen, k. heljanko, i. niemelä. planning as satisfiability: parallel plans and algorithms for plan search. artif intell 170(12-13):1031–1080, 2006. doi:10.1016/j.artint.2006.08.002. [4] n. robinson, c. gretton, d. n. pham, a. sattar. sat-based parallel planning using a split representation of actions. in a. gerevini, a. e. howe, a. cesta, i. refanidis (eds.), icaps. aaai, 2009. http://aaai. org/ocs/index.php/icaps/icaps09/paper/view/732. [5] t. balyo. relaxing the relaxed exist-step parallel planning semantics. in ictai, pp. 865–871. ieee, 2013. doi:10.1109/ictai.2013.131. [6] j. rintanen. planning as satisfiability: heuristics. artif intell 193:45–86, 2012. doi:10.1016/j.artint.2012.08.001. [7] c. l. lópez, s. j. celorrio, á. g. olaya. the deterministic part of the seventh international planning competition. artificial intelligence 2015. doi:10.1016/j.artint.2015.01.004. [8] c. bäckström, b. nebel. complexity results for sas+ planning. computational intelligence 11:625–656, 1995. doi:10.1111/j.1467-8640.1995.tb00052.x. [9] r. fikes, n. j. nilsson. strips: a new approach to the application of theorem proving to problem solving. artif intell 2(3/4):189–208, 1971. doi:10.1016/0004-3702(71)90010-5. [10] a. blum, m. l. furst. fast planning through planning graph analysis. artificial intelligence 90(12):281–300, 1997. doi:10.1016/s0004-3702(96)00047-1. [11] a. biere. lingeling and plingeling home page, 2015. http://fmv.jku.at/lingeling/. [12] m. helmert. the fast downward planning system. journal of artificial intelligence research (jair) 26:191–246, 2006. doi:10.1613/jair.1705. 7 http://citeseerx.ist.psu.edu/viewdoc/summary?doi=10.1.1.35.9443 http://citeseerx.ist.psu.edu/viewdoc/summary?doi=10.1.1.35.9443 http://citeseerx.ist.psu.edu/viewdoc/summary?doi=10.1.1.182.9561 http://citeseerx.ist.psu.edu/viewdoc/summary?doi=10.1.1.182.9561 http://dx.doi.org/10.1016/j.artint.2006.08.002 http://aaai.org/ocs/index.php/icaps/icaps09/paper/view/732 http://aaai.org/ocs/index.php/icaps/icaps09/paper/view/732 http://dx.doi.org/10.1109/ictai.2013.131 http://dx.doi.org/10.1016/j.artint.2012.08.001 http://dx.doi.org/10.1016/j.artint.2015.01.004 http://dx.doi.org/10.1111/j.1467-8640.1995.tb00052.x http://dx.doi.org/10.1016/0004-3702(71)90010-5 http://dx.doi.org/10.1016/s0004-3702(96)00047-1 http://fmv.jku.at/lingeling/ http://dx.doi.org/10.1613/jair.1705 acta polytechnica ctu proceedings 2:1–7, 2015 1 introduction 2 preliminaries 2.1 satisfiability 2.2 planning 2.3 parallel plans 3 finding plans using sat 4 reinforced encoding 4.1 correctness 4.2 size of the encoded formula 5 experimental evaluation 5.1 experimental setting 5.2 experimental results 6 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0010 acta polytechnica ctu proceedings 12:10–12, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app qfd example in interaction with hmi inga chudjakova∗, jaromír tobiška škoda auto a.s., tř. václava klementa 869, mladá boleslav, 293 60, czech republic ∗ corresponding author: inga.kadlecova@skoda-auto.cz abstract. this article focuses on the application of the quality function deployment (qfd) method and shows how we can use this method in the automotive industry. the cockpit of the modern car is still developing and changing according to technical progress and customers’ requirements. the qfd method enables the setting of the customers’ requirements and then puts them into technical expression. the matrix of qfd sets those technical expressions that are the most important properties customers expect. these chosen properties can be tested during the technical development. the testing of new hmi concepts goes through a driving simulator and can suggest which one is suitable for being used in a real car. keywords: the qfd method, customers’ requirements, automotive industry, doe, hmi, driving simulator, house of quality. 1. introduction nowadays in the automotive industry, the aim of all car producers is customer satisfaction, but also technical progress, that can differentiate them in the current competitive environment. a set of quality methods serves/is used to achieve customer satisfaction in the automotive industry. they lead to production of high-quality production and are also related to higher profit. 2. quality assurance methods for the reasons above it is necessary to use the quality assurance methods in the automotive industry and use them at the beginning of the product development process. the qfd method, i.e. quality function deployment, is one of these methods. this method is used mostly at the first stage of the product development process because the qfd method allows transformation of customer requirements into technical expression[1]. its main task is the determination of customer requirements and their implementation in the process of the product preparation. thanks to qfd methods, the designers can understand what customers want and it designs the product. the methods used for optimizing the costs and the methods to prevent errors and faults in the development of a product also belong among the assurance methods. for more detailed classification of the quality assurance methods overview see the fig. 1. by using qfd method(s) we can assess the fulfilment of the customers’ requirement for products and processes. at the start of the qfd method application, there are the customer’s requirements (wishes). to find out what the customers want we need to focus on the following: a)technical parameters, a physical element, a specified part of system figure 1. the quality assurance methods overview. b)defined procecesses, assembly operations, handling operations, specific actions figure 2. the qfd application. the use of the qfd method is equal in the both cases. the difference is in the outcomes. we are able to find the key element of the system and the specific technical parameter, which we monitor to ensure the fulfilment of customers’ requirements or key activities; the process which implementation may impact the consequences effecting the customers’ satisfaction. 10 http://dx.doi.org/10.14311/app.2017.12.0010 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 qfd example in interaction with hmi 3. the application of the qfd method the basic outcomes of the qfd (quality function deployment) method application are the answers to the two following questions: • what are the customer’s expectations? • how can we meet these expectations? [1] the qfd method is characterized by four application stages, which are as follows: product proposal (a proposal of characteristic product properties), proposal of components, processes planning and production planning. each of these stages is noted in a diagram and matrixes. 3.1. matrixes of the qfd method the qfd method is based on the qfd matrix, which is often marked as the "house of quality" and consists of a team of experts. the overall structure of the qfd matrix has 8 specified fields, but during the use the qfd matrix not all fields need to be completed. it is possible to use only a part of the qfd matrix, as required by the project. the qfd matrix that describes the most important field, can be seen at the graph bellow (see fig. 3). figure 3. qfd matrix "the house of quality". for the completing of the qfd matrix it is not required to use a special programme or software, one can just fill an excel table. the most important thing for the using the qfd method and filling in the qfd matrix is to know the succession of inputs and outputs – for detail see fig. 4. the qfd method can interpret the customers’ requirements, which are often defined only verbally, into the technical values that the designers can work with. to use the qfd method correctly, it is essential to get the inputs. this information can be obtained from: • surveys/questionnaires • thinking loud (lautes denken) figure 4. qfd method relations. • clinical studies • group discussions/think tanks ??? • j. d. power these procedures are established in quantitative and qualitative research techniques. in order to receive the information needed, it is necessary to choose the right method, group of respondents, questionnaire, and also the form of responses for each project. the application of the entire qfd method is based on obtaining these input data. then the matrix is completed with technical solutions that could have an influence on the customers’ requirements. the technical parameters may be defined by using accurate parameters or units. therefore, the best technical solutions are those that influence several customer wishes at once. the correlation between the customers’ expectations and the technical parameters is placed in the main field of the qfd matrix. this mutual correlation shows how much the technical solution will influence the customers’ expectations. the mutual influence may have the following values: 0 = no influence, 1 = low influence, 3 = average influence, or 9 = huge influence. 4. a practical example of application of the qfd method in the automotive industry at the present time, the load of the driver while driving is increasing. this is caused by the improvement of technology and development of the hmi systems in cars. also, the proportion of displaying and controlling components in the cockpit is increasing. an interaction between these car systems, displaying components and a driver may affect driving safety in a remarkable way[2]. for the practical example of using qfd method, we need to optimize the display to ensure safe operation while driving. for setting of the car display properties of we used qfd method. in order to use the qfd method, firstly we needed to find out which display properties customers found the most important. these were the inputs into the qfd matrix and we had used the qualitative and quantitative techniques and methods. the input data were received due to an evaluation of the questionnaires, 11 inga chudjakova, jaromír tobiška acta polytechnica ctu proceedings where respondents assessed the general properties of a display independent of the tested concepts. for these customers’ expectations, the technical experts set the technical parameters, which is are in the field p4. then the teams set the correlations between the customers’ expectations and technical parameters at the field p5. after the calculations of qfd matrix, in the field p6 we can see which technical parameters had effected the customers’ expectations the most. figure 5. example of qfd matrix. the highest rated technical expressions from the qfd matrix were tested in the technical clinic through a driving simulator. in these clinics, the highest rated expressions were: a font size, font legibility, text understanding and contrast. these parameters were tested separately on a separate screen. several chosen screens from the display menu were tested. they display various combinations of the script sizes and contrast. the respondents were expected to assess the individual items on each screen, such as a script legibility, text comprehension and contrast. the above-mentioned testing revealed that the best combination of good display legibility for customers is the second biggest letter size of the tested script with combination of the best contrast (contrast 1), because the biggest font size can be distracting[3]. figure 6. combination of fontsize and contrast. 5. conclusion the qfd method is an integral part of a product development in the automobile industry. its contribution is provable in all the stages of the car development process as well as in the area of the hmi concept testing. the main objective of the hmi systems testing is the minimization of the time that these systems are controlled and thus the time that the driver is not concentrating on driving. together, these methods contribute to customers’ satisfaction, which has been nowadays emphasized more and more in the automotive industry. references [1] j. machan, j. tobiška, et al. metody kvality užívané ve fázi vývoje výrobku – aplikace v automobilovém průmyslu. ii.a revised and expanded edition, prague 2012, 117 p., isbn 978-80-87042-50-2. [2] m. novák, p. bouchner, j. faber, et al. senioři za volantem. prague: faculty of transportation sciences, czech technical university in prague, 2008, isbn 978-80-87136-20-7. [3] zvláštní projekty elektrostrategie a výzkumu, technický vývoj, škoda auto a.s.: projekt hmi ve škoda auto. (prezentace), mladá boleslav, an internal document, 2013. 12 acta polytechnica ctu proceedings 12:10–12, 2017 1 introduction 2 quality assurance methods 3 the application of the qfd method 3.1 matrixes of the qfd method 4 a practical example of application of the qfd method in the automotive industry 5 conclusion references 181 acta polytechnica ctu proceedings 1(1): 181–188, 2014 181 doi: 10.14311/app.2014.01.0181 magnetorotational explosions of core-collapse supernovae gennady s. bisnovatyi-kogan1, sergey g. moiseenko2, nikolay v. ardeljan3 1space research institute, profsoyuznaya str. 84/32, moscow 117997, russia, and national research nuclear university ”mephi”, kashirskoye shosse, 31, moscow 115409, russia 2space research institute, profsoyuznaya str. 84/32, moscow 117997, russia 3department of computational mathematics and cybernetics, moscow state university, vorobjevy gory, moscow b-234, russia corresponding author: gkogan@iki.rssi.ru abstract core-collapse supernovae are accompanied by formation of neutron stars. the gravitation energy is transformed into the energy of the explosion, observed as sn ii, sn ib,c type supernovae. we present results of 2-d mhd simulations, where the source of energy is rotation, and magnetic field serves as a ”transition belt” for the transformation of the rotation energy into the energy of the explosion. the toroidal part of the magnetic energy initially grows linearly with time due to differential rotation. when the twisted toroidal component strongly exceeds the poloidal field, magneto-rotational instability develops, leading to a drastic acceleration in the growth of magnetic energy. finally, a fast mhd shock is formed, producing a supernova explosion. mildly collimated jet is produced for dipole-like type of the initial field. at very high initial magnetic field no mri development was found. keywords: core-collapse supernova magnetorotational mechanism numerical modeling. 1 introduction supernova is one of the most powerful explosion in the universe which releases about 1051 erg both in radiation and kinetic energies. sne explode at the end of evolution of massive stars, with initial mass larger than ∼ 8m�. a thermonuclear explosion of c-o degenerate core with total disruption of the star takes place in sn ia, what happens when initial mass of the star does not exceed ∼ 12m�, when electrons are degenerate in the carbon-oxygen core . for larger initial masses the evolution proceeds until the formation of the iron core, and the star collapses due to a loss of a hydrodynamic stability. during the core collapse and formation of a neutron star, gravitational energy release ∼ 6 · 1053 erg, is carried away by neutrino. the first mechanism suggested in [15] for the explanation of the explosion in a corecollapse sn was connected with a neutrino deposition. the huge energy flux carried by neutrino heats the infalling outer layers, reverse the direction of motion, and leads to formation of the shock wave, producing explosions of sn ii, sn ib,c types. later more accurate calculations revealed that the energy of such explosion is not enough for the explanation of observations. many modifications of the neutrino model have been calculated during years, but the problem is not yet clear. the review of the problem may be found in the book [8]. in recent simulations (see details in [23]) we have found that for extremely strong initial magnetic field h0 = 10 12g a prompt supernova explosion occurs, and collimated jet is formed in agreement with [28]. for case when the initial magnetic field is weaker h0 = 10 9g we have identified, after the linear growth of the poloidal magnetic field due to differential rotation, the exponential field growth due to the magnetorotational instability of tayler type [30]. we call the combination of differential rotation with the tayler type mri instability, as magneto-differential-rotation instability (mdri). the ejection due to the explosion is only weakly collimated, while [28] had obtained a strong collimation in this variant also. in these simulations we considered a uniform magnetic field along the rotational axis, as the initial field configuration, similar to [28]. 2 magnetorotational mechanism of explosion in magnetorotational explosion (mre) the transformation of the rotational energy of the neutron star into explosion energy takes place by means of the magnetic field [7]). neutron stars are rotating, and have magnetic fields up to 1013 gs, and even more. often one or two-side ejections are visible. that indicate to nonspherical form of the sne explosions. in differentially 181 http://dx.doi.org/10.14311/app.2014.01.0181 gennady s. bisnovatyi-kogan, sergey g. moiseenko, nikolay v. ardeljan rotating new born neutron stars radial magnetic field is twisted, and magnetic pressure becomes very high, producing mhd shock by which the rotational energy is transformed to the explosion energy. calculations of mre have been done in [12], using one-dimensional nonstationary equations of magnetic hydrodynamics, for the case of cylindrical symmetry. the energy source is supposed to be the rotational energy of the system (the neutron star, and surrounding envelope). the calculations show that the envelope splits up during the dynamical evolution of the system, the main part of the envelope joins the neutron star and becomes uniformly rotating with it, and the outer part of the envelope expands with large velocity, carrying out a considerable part of rotational energy and rotational momentum. mre has an efficiency about 10% of the rotational energy, the ejected mass is ≈ 0.1 of the star mass, explosion energy ≈ 1051 erg. ejected mass and explosion energy depend weekly on the parameter α = emag/egrav at initial moment. explosion time depends on α as texpl ∼ 1√α. small α is difficult for numerical calculations with explicit numerical schemes because of the courant restriction on the time step, hard system of equations, where α determines a hardness. 3 2-d calculations the numerical method used in simulations is based on the implicit operator-difference, completely conservative scheme on a lagrangian triangular grid of variable structure, with grid reconstruction (fig.1). the implicitness of the applied numerical scheme allows for large time-steps. it is important to use the implicit scheme in the presence of two strongly different time-scales: the small one due to huge sound velocity in the central parts of the star, and the big one determining the evolution of the magnetic field. the method applied here was developed and its stability was investigated in the papers of [5], [6], [1]. the scheme is fully conservative, what includes conservation of mass, momentum and total energy, and correct transitions between different types of energies. it was tested thoroughly with different tests by [2]. in the calculations of magnetorotational corecollapse supernova performed by [3], magnetohydrodynamic (mhd) equations with self-gravitation, and infinite conductivity have been solved using the numerical scheme as described above. the problem has an axial symmetry ( ∂ ∂φ = 0), and the symmetry to the equatorial plane (z=0). initial toroidal current jφ was taken at the initial moment (time started now from the stationary rotating neutron star) producing hr, hz according to biot-savart law b = 1 c ∫ v j×r r3 dv . initial magnetic field of quadrupole-like symmetry is obtained at opposite directions of the current in both hemispheres. neutrino cooling was calculated using a variant of a fluxlimited method, [3]. figure 1: example of the triangular grid magnetic field is amplified due to twisting by the differential rotation, and subsequent development of the magnetorotational instability. the field distribution for initial quadrupole-like magnetic field with α = 10−6, at the moment of the maximal energy of the toroidal magnetic field is represented in fig.1. the maximal value of bφ = 2.5·1016 gs was obtained in the calculations. the magnetic field at the surface of the neutron star after the explosion is b = 4 ·1012 gs. time dependence during the explosion of rotational, gravitational, internal, and kinetic poloidal energies is given in figs.3. almost all gravitational energy, transforming into heat during the collapse, is carried away by weakly interacting neutrino. the total energy ejected in the kinetic form is ∼ 0.6 · 1051 erg, and the total ejected mass is equal to ∼ 0.14m�. figure 2: toroidal magnetic field distribution at the moment of its maximal energy for the initial quadrupole field . 182 magnetorotational explosions of core-collapse supernovae time, s 0 0.1 0.2 1 2 3 4 erot[10 51 ergs] ekinpol[10 51 ergs] figure 3: time dependence of rotational, kinetic poloidal, and magnetic energies during explosion for a dipole -like field, from [22]. the simulations were done for the initial poloidal magnetic field of quadrupole [3] and of dipole [22] types of symmetry. before the collapse the ratios between the rotational and gravitational, and between the internal and gravitational energies of the star had been chosen as: erot egrav = 0.0057, eint egrav = 0.727. the initial magnetic field was ”turned on” after the collapse stage. the ratio between the initial magnetic and gravitational energies was chosen as 10−6. the initial poloidal magnetic field in the center, at start of the evolution of the toroidal field was ∼ 3.2 × 1013g. time, s e je ct e d m a ss /m su n 0 0.1 0.2 0.30.00 0.04 0.08 0.12 figure 4: time dependence of the ejected mass during the magnetorotational explosion with initial dipole magnetic field, from [22]. the magnetic field works as a piston for the originated mhd shock. the time dependence of the ejected mass and energy is given in figs. 4, 5. during the magnetorotational explosion ∼ 0.14 m� of the mass and ∼ 0.6 · 1051ergs (∼ 10% of the rotational energy) are ejected. the simulation of the mr supernova explosion for various initial core masses and rotational energies was done by [11]. the initial core mass was varied from 1.2m� to 1.7m�, the initial specific rotational energy erot/mcore, was varied from 0.19 × 1019 to 0.4 × 1019 erg/g. the explosive energy increases with the mass of the core, and the initial rotational energy. the energy released in mr explosion, (0.5−2.6)×1051 erg, is sufficient to explain supernova with collapsing cores, types ii and ib. the energies of type ic supernovae could be higher. time, s e je ct e d e n e rg y [1 0 5 1 e rg s] 0.2 0.3 0.1 0.2 0.3 0.4 0.5 0.6 figure 5: time dependence of the ejected energy during the magnetorotational explosion with initial dipole magnetic field, from [22]. 4 magnetorotational instability magnetorotational instability (mri) leads to exponential growth of magnetic fields. different types of mri have been studied by [31], [27]. mri starts to develop when the ratio of the toroidal to poloidal magnetic energies is becoming large. in 1-d calculations mri is absent because of a restricted degree of freedom, and time of mr explosion is increasing with α as texpl ∼ 1√α, α = emag0 egrav0 . due to development of mri the time of mr explosion depends on α much weaker. the mr explosion happens when the magnetic energy is becoming comparable to the internal energy, at least in some parts of the star. while the starting magnetic energy linearly depends on α, and mri leads to exponential growth of the magnetic energy, the total time of mre in 2-d is growing logarithmically with decreasing of α, texpl ∼− log α. these dependencies are seen clearly from 1-d ([12]) and 2-d calculations ([3], [22]) giving the following explosion times texpl (in arbitrary units): α = 0.01, texpl = 10, α = 10 −12, texpl = 10 6 in 1d, and α = 10−6, texpl ∼ 6, α = 10−12, texpl ∼ 12 in 2-d. the dependence of the explosion time is shown 183 gennady s. bisnovatyi-kogan, sergey g. moiseenko, nikolay v. ardeljan in graphs for the quadrupole ([21]), and dipole ([22]) configurations of the magnetic field. the qualitative picture of the mri in 2d, and the example of the analytical toy model with an exponential growth of the magnetic field, have been presented by [3], [22]. 5 jet formation in mre jet formation in mre happens when the initial magnetic field is of a dipole-like structure. 2-d calculations with the initial dipole-like magnetic field gave almost the same values of the energy of explosion ∼ 0.5 · 1051 , and ejected mass ≈ 0.14m�, but the outburst was slightly collimated along the rotational axis [22], see fig.6 and fig.7. r, km z, km 0 2000 4000 6000 0 1000 2000 3000 4000 5000 6000 7000 0 500 10000 200 400 600 800 1000 figure 6: time evolution of the velocity field (outflow) for the time moment t = 0.075s, from [22] 0 500 10000 200 400 600 800 1000 r, km z, km 0 2000 4000 6000 0 1000 2000 3000 4000 5000 6000 7000 figure 7: time evolution of the velocity field (outflow) for the time moment, t = 0.25s, from [22]. simulations of the mr supernova have been made with equation of state suggested in [25]. a comparison of our results for the initially uniform magnetic field, using a lagrangian scheme, with the results in [28] and [29], using an eulerian scheme for the same initial and boundary conditions, shows good agreement for a strong initial field (h0 = 10 12g), while for a weaker field (h0 = 10 9g) we get mildly collimated jet-like explosion; see also [13]. details of the results of these simulations will be published elsewhere ([23]). mri is developed in the case of a weaker initial magnetic field, and it is not present in the calculations with stronger field, see fig.8 and fig.12. we have made simulations for the initial magnetic fields h0 = 10 9, 1012g. the ratio of the initial rotational energy to the absolute value of the gravitational energy was taken erot0/egrav0 = 1%, 2%. when the initial magnetic field is moderate (h0 = 10 9 g) and erot0/egrav0 = 1% mdri develops what means exponential growth of all components of the magnetic field (fig.8). figure 8: developed mdri due to convection and mri/tayler instability at t = 267ms for the case h0 = 109g,erot0/egrav0 = 1% (contour plot the toroidal magnetic field, arrow lines force lines of the poloidal magnetic field). at fig.9 the lagrangian triangular grid and the ratio of the toroidal magnetic energy to the poloidal one (etor/epol) is represented for the case h0 = 10 9 g, and erot0/egrav0 = 1%. the toroidal magnetic energy dominates over the poloidal one in the significant part of the region where new neutron star is forming. the mdri is well-resolved on our triangular grid. in the case (h0 = 10 12 g) and erot0/egrav0 = 1% there is no regions of domination of etor over epol. the fig.10 represents a time evolution of rotational, magnetic poloidal and toroidal energies for mr explo184 magnetorotational explosions of core-collapse supernovae sion when h0 = 10 9g. the fig.11 is the same data plot as the fig.10 but zoomed and the vertical axis is in logarithmic scale. the straight dash-dotted line at the fig.11 shows the exponential growth of the toroidal and poloidal magnetic energies with the time due to the mdri. the rotational energy has two maxima. the first contraction is accompanied by the strong growth of the rotational energy due to angular momentum conservation, maximum of which coincides with the first maximum of the density. the first contraction, and the subsequent bounce, happens when the magnetic field is growing slowly, and the angular momentum losses from the stellar core are small. development of the magnetorotational instability leads to a rapid growth of the magnetic field, large angular momentum flux from the core, what stops the expansion, and leads to the second contraction phase. in this case the contraction is not transforms into expansion, because of the rapid decrease of the rotational energy due to strong angular momentum flux outside from the core. r,cm z ,c m 0 5e+060 2e+06 4e+06 6e+06 eetor/eepol 1000 800 600 500 200 100 10 1 time= 0.265sec figure 9: the lagrangian triangular grid and the ratio of the toroidal magnetic energy etor to the poloidal magnetic energy epol, etor epol at t = 265ms for the case h0 = 10 9g,erot0/egrav0 = 1%. the energy of the poloidal magnetic field grows due to the contraction until the time t ≈ 0.225 sec. then it slightly decreases because of the formation of the bounce shock, and its motion outwards. the toroidal magnetic energy grows as quadratic function because of wrapping of the magnetic force lines (toroidal component of the magnetic field grows linearly). starting from t ≈ 0.3 sec both the poloidal and toroidal magnetic energies begin to grow exponentially due to mdri. at t ≈ 0.36 sec both magnetic energies comes to saturation. the mhd shock wave develops what leads to the mr explosion. the mr explosion develops in all directions without formation of a collimated flow. the mr explosion for an extremely high initial magnetic field (h0 = 10 12 g) is developing in a qualitatively different way. the initial magnetic field is so strong that it grows strongly during the first contraction, and the explosion happens before the development of mdri happens (fig.12). at the fig.13 the time evolution of the rotational, poloidal magnetic and toroidal magnetic energies are represented. the rotational energy has one extremum at t ≈ 0.32 sec corresponds to the maximal contraction, accompanied by a corresponding growth of of the toroidal and poloidal magnetic energies. the poloidal magnetic field grows due to the contraction, the toroidal magnetic field appears due to the differential rotation and is amplified both due to the differential rotation and the contraction of the core. the strong initial magnetic field leads to a rapid loss of the angular momentum from the core already during the contraction phase. the centrifugal force becomes unimportant, and the first contraction is not followed by any bounce, leaving behind a dense slowly rotating neutron star core. we have got here a prompt explosion. the force lines of the magnetic field play the role of ’rails’. the matter moves along the force lines. the magnetic pressure dominates a the periphery of the core. the mr explosion develops mainly along the axis of rotation, and the collimated flow (protojet) is formed. the mr explosion results in the collimated jet. the degree of jet collimation is approximately the same as in [28]. for the case when erot0/egrav0 = 1% and h0 = 10 9 g the mr supernova explosion energy is ∼ 4 × 1050 erg, for the h0 = 1012 g the mr supernova explosion energy reaches the value of ∼ 8 × 1050erg. the explosion energy resulted in the simulations by lagrangian method are close to those ones found in the simulations made by eulerian scheme [28] (excluding the case of b0 = 10 12 g and erot0/|egrav0| = 1/%). 6 asymmetry of the explosion it is known from the observations that the shapes of core collapse supernovae are different. from our simulations it follows that mr supernova explosion arises after development of the mri. the development of the mri is a stochastic process and hence the resulting shape of the supernova can vary. we may conclude that mr supernova explosion mechanism can lead to different shape of the supernova. it is important to point out that mr mechanism of supernova explosion leads always to asymmetrical outbursts. the simulations of the mr supernova explosions described here are restricted by the symmetry to the equa185 gennady s. bisnovatyi-kogan, sergey g. moiseenko, nikolay v. ardeljan torial plane. while in reality this symmetry can be violated due to the mri, simultaneous presence of the initial dipole and quadrupole -like magnetic field ([33]) and initial toroidal magnetic field ([10]). the violation of the symmetry could lead to the kick effect and formation of rapidly moving radio pulsars. a kick velocity, along the rotational axis, formed due to magnetohydrodynamic processes in presence of the asymmetry of the magnetic field, by estimations [9] does not exceed 300km/sec. figure 10: time evolution of rotational energy erot (solid line), magnetic poloidal energy emagpol (dashed line) and magnetic toroidal energy emagtor (dashdotted line) for the case h0 = 10 9g,erot0/egrav0 = 1%. when rotational and magnetic axes do not coincide the whole picture of the explosion process is three dimensional. nevertheless, the magnetic field twisting happens always around the rotational axis, so we may expect the kick velocity of the neutron star be strongly correlated with its spin direction. during the phase of mre explosion the regular component of magnetic field may exceed temporally 1016 g [3], [22], when the neutrino cross-section depends on the magnetic field strength. the level of the anisotropy of the magnetic field relative to the plain perpendicular to the rotational axis [20] may be of the order of 50%, leading to strong anisotropy of the neutrino flux. the kick velocity due to the anisotropy of the neutrino flux may reach several thousands km/c [9], explaining appearance of the most rapidly moving radio pulsars [32]. simultaneously, because of the stochastic nature of mri, the level of the anisotropy should be strongly variable, leading to a large spreading in the the neutron star velocities. this prediction of mr explosion differs from the models with a powerful neutrino convection, where arbitrary direction of the kick velocity is expected ([14],[17]). it was claimed in [16], that proto-neutron star (pns) convection was found to be a secondary feature of the corecollapse phenomenon, rather than a decisive ingredient for a successful explosion. figure 11: zoomed time evolution of rotational energy erot (solid line), magnetic poloidal energy emagpol (dashed line) and magnetic toroidal energy emagtor (dash-dotted line) for the case h0 = 109g,erot0/egrav0 = 1%. straight dash-dotted line shows exponential growth of the toroidal and poloidal magnetic energies. analysis of observations of pulsars shows that rotation and velocity vectors of pulsars are aligned, as is predicted by the mr supernova mechanism. this alignment was first found in [26], and was confirmed with reliability, increasing with time, in the papers [18],[19],[24]. the alignment of the vectors can be violated in the case when the supernova explodes in a binary system. 7 conclusions in mre the efficiency of transformation of rotational energy into the energy of explosion is ∼ 10%. mri strongly accelerates mre, at lower values of the initial magnetic fields. jet formation is possible for dipolelike topology of the field. mre energy is not sensitive to the details of the equation of state, model of the neutrino transfer, and to the choice of the numerical scheme. the observed alignment of the rotation and velocity vectors of pulsars follows directly from the mre supernova model. 186 magnetorotational explosions of core-collapse supernovae r,cm z ,c m 0 1e+06 2e+06 0.0e+00 5.0e+05 1.0e+06 1.5e+06 2.0e+06 time= 0.328sec figure 12: absence of mdri at t = 328ms for the case h0 = 10 12g,erot0/egrav0 = 1% (contour plot the toroidal magnetic field, arrow lines force lines of the poloidal magnetic field). 0.1 0.15 0.2 0.25 0.3 0.35 time, sec e rot e magpol e magtor 10 8 6 4 2 e n e rg y x 1 0 5 0 , e rg figure 13: time evolution of rotational energy erot (solid line), magnetic poloidal energy emagpol (dashed line) and magnetic toroidal energy emagtor (dashdotted line) for the case h0 = 10 12g,erot0/egrav0 = 1%. acknowledgement the work of sgm and gsbk was supported partially by rfbr grant 11-02-00602, grant for leading scientific schools nsh-5440.2012.2 and program of ras ’origin structure and evolution of objects in the universe’. references [1] ardeljan, n.v., bisnovatyi-kogan, g.s., kosmachevskii k.v. & moiseenko, s.g. 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[28] takiwaki, t., kotake, k., nagataki, s., & sato, k. 2004, apj, 616, 1086 doi:10.1086/424993 [29] takiwaki, t., kotake, k., & sato, k. 2009, apj, 691, 1360 doi:10.1088/0004-637x/691/2/1360 [30] tayler, r. 1973, mnras, 161, 365 [31] velikhov, e.p. 1959, j. exper. theor. phys., 36, 1398 [32] vlemmings, w.h.t., et al. 2005, mem. sai, 2005, 76, 531 (astro-ph/0509025) [33] wang, j. c. l., sulkanen, m. e., lovelace, r. v. e., 1992, apj, 390, 46 188 http://dx.doi.org/10.1086/589759 http://dx.doi.org/10.1111/j.1365-2966.2006.10517.x http://dx.doi.org/10.1111/j.1365-2966.2012.21083.x http://dx.doi.org/10.1086/424993 http://dx.doi.org/10.1088/0004-637x/691/2/1360 introduction magnetorotational mechanism of explosion 2-d calculations magnetorotational instability jet formation in mre asymmetry of the explosion conclusions 56 acta polytechnica ctu proceedings 1(1): 56–65, 2014 56 doi: 10.14311/app.2014.01.0056 multi-frequency study of the sz effect in cosmic structures sergio colafrancesco1 1university of the witwatersrand, johannesburg (south africa) corresponding author: sergio colafrancesco. email: sergio.colafrancesco@wits.ac.za abstract the sunyaev-zel’dovich effect (sze) is a relevant probe for cosmology and astrophysics. a multi-frequency approach to study the sze in cosmic structures turns out to be crucial in the use of this probe for astrophysics and cosmology. astrophysical and cosmological applications to galaxy clusters, galaxies, radiogalaxies and large-scale structures are discussed. future directions for the study of the sze and its polarization are finally outlined. keywords: cosmology cmb dark matter dark energy cosmic magnetism cosmic structures: galaxy clusters, galaxies, radio galaxies. 1 introduction galaxy clusters, the largest gravitationally bound structures in the universe, are the representative systems of the distribution of lss in the universe. the description of these cosmic structures is continuously enriching of physical details regarding their matter and field content. dark matter (dm) is the dominant form of matter that creates the potential wells of cosmic structures, from those on the largest scales down to galactic and sub-galactic scales. if we consistently take into account the fundamental nature of dm particles, we are inevitably bound to consider the effects of their annihilation or decay on the structure and evolution of the dm clumps [13]. baryonic material collected in the dm potential wells of lss is likely heated and shocked by large–scale shock waves found in cosmological simulations (see, e.g, [53]) which produce a complex distribution of mach numbers. the presence of shock waves with relatively high mach numbers naively suggest that fermi-like acceleration might take place in lss thereby accelerating cosmic rays (crs) that can be efficiently confined mainly in galaxy clusters [8], while they can diffuse out of galactic and sub-galactic structures [8, 9]. radio observations and magneto-hydrodinamic (mhd) simulations also suggest that magnetic fields are associated to the distribution of baryonic material collected in lss potential wells. the seed magnetic field is likely amplified and made turbulent by the coupling of gravitational collapse and mhd processes during the formation of galaxy clusters. evidence for wide-scale and turbulent intra-cluster magnetic field is indicated by the diffuse radio synchrotron emission found in many clusters (with typical values b ∼ 0.1 − 2 µg), by radio relics in dynamically active clusters (with typical values b ∼ 0.2 − 5 µg) and by faraday rotation measurements of background and embedded polarized radio sources (with typical values b ∼ 1 − 50 µg [27]. very massive dm clumps that collapse at high redshifts (z ∼ 6−7) often contain the most massive black holes (bhs) at their centers [35]. the agn descendants of these ancient supermassive bhs (smbhs) are found to be part of the massive galaxies located at the centers of the most massive galaxy clusters we observe at the present epoch, like, e.g., cena, m87/virgo, perseus, a262, a4059, with their radio lobes penetrating the icm for tens or hundreds of kpcs. it is often observed that the radio jets/lobes end up in approximately spherical bubbles of relativistic plasma (likely containing relativistic or mildly-relativistic plasmas) that appear as cavities in the x-ray images of galaxy clusters [3] with dimensions ranging from a few kpcs (as in perseus) to ∼ 100 kpc. the combination of high-resolution radio and x-ray images indicates that the relativistic plasma found inside the x-ray cavities is connected with the jet/lobe structure of the central agn and with the history of the ejection events and mechanisms from the central agn. the cluster cores which host such non-thermal phenomena (bhs, cavities, magnetic fields) are found to be systematically cooler than the outer regions of the clusters with the inner temperature setting at a value ∼ 1 3 − 1 2 of the outer temperature, usually consistent with the virial expectation. it seems that a heating agent of nongravitational and non-thermal nature – with a heating rate that is able to accomodate itself to the cooling rate 56 http://dx.doi.org/10.14311/app.2014.01.0056 multi-frequency study of the sz effect in cosmic structures of the intra-cluster plasma – is present in the cluster’s cool core so as to maintain it in a quasi-stationary, warm configuration [16]. the previous evidence indicates that galaxy clusters are the largest storage rooms for cosmic material (galaxies, dm, hot thermal baryonic plasma, non-thermal and relativistic plasma, bhs, magnetic fields, crs). in this sense they can be considered as the largest multidisciplinary laboratories in the universe where one can efficiently study some of the most interesting aspects of the astrophysics of lss: the nature of dm, the origin and distribution of crs, the impact of magnetic fields on lss, the impact of bhs on lss, the interplay between thermal and non-thermal phenomena in lss. two ways can be identified to proceed in this study: i) a multi-technique, single-purpose approach that require to integrate the analysis of various observations performed in various frequency bands; ii) a singletechnique, multipurpose approach that is able to provide detailed physical information based on a single observational technique. we will discuss, in the following, some of the results obtainable through the last approach by using the physical information contained into the sze. 2 the physics of the sz effect the sze is produced by the inverse compton scattering (ics) of cmb photons off the electrons confined in the atmospheres of cosmic structures. manifestations of the sze include: i) spectral distortions of the cmb due to up-scattering of cmb photons induced by high-e electrons (thermal, non-thermal and relativistic sze); ii) spectral distortion of the cmb due to a bulk motion of the electronic plasma w.r.t. the hubble flow (kinematic sze); iii) polarization of the cmb due to dynamical and plasma effects (sze polarization). the sz effect. the spectral distortion of the cmb spectrum observable in the direction of a galaxy cluster writes [56, 2, 6] as ∆i(x) = 2 (ktcmb) 3 (hc)2 y g(x) , (1) where ∆i(x) = i(x) − i0(x), i(x) is the up-scattered cmb spectrum in the direction of the cluster and i0(x) is the unscattered cmb spectrum in the direction of a sky area contiguous to the cluster. here x ≡ hν/ktcmb, h is the planck constant, k is the boltzmann constant, tcmb = 2.726 k is the cmb temperature today and ν is the observing frequency. the comptonization parameter y is y = σt mec2 ∫ ped` (2) in terms of the pressure pe contributed by the electronic population. here σt is the thomson cross section, me the electron mass, and c the speed of light. the spectral function g(x) of the sze is g(x) = mec 2 〈εe〉 { 1 τe [∫ +∞ −∞ i0(xe −s)p(s)ds− i0(x) ]} (3) [6] in terms of the photon redistribution function p(s) and of i0(x) = i0(x)/[2(ktcmb) 3/(hc)2] = x3/(ex − 1) . (4) the quantity 〈εe〉≡ σt τe ∫ ped` = ∫ ∞ 0 dpfe(p) 1 3 pv(p)mec , (5) where fe(p) is the normalized electron momentum distribution function, is the average energy of the electron plasma [6]. the optical depth along the line of sight ` of the electron population with number density ne is τe = σt ∫ d`ne . (6) the photon redistribution function p(s), with s = ln(ν′/ν) in terms of the cmb photon frequency increase factor ν′/ν, can be calculated by repeated convolution of the single-scattering redistribution function, p1(s) = ∫ dpfe(p)ps(s; p), where ps(s; p) depends on the physics of inverse compton scattering. the previous description is relativistically covariant and general enough to be applied to both thermal and nonthermal plasma, as well as to a combination of the two (see [6, 17, 38] for details). kinematic sz effect. the velocity (or kinematic) sze (hereafter ksze) arises if the plasma causing the thermal, or non-thermal, sze is moving relative to the hubble flow. in the reference frame of the scattering particle the cmb radiation appears anisotropic, and the effect of the ics is to re-isotropize the radiation slightly. back in the rest frame of the observer the radiation field is no longer isotropic, but shows a structure towards the scattering atmosphere with amplitude ∝ τevt/c, where vt is the component of the peculiar velocity of the scattering atmosphere along the line of sight [55, 51]. the brightness change of the cmb due to the ksze is given by ∆i i = −τeβt xex ex − 1 (7) with βt ≡ vtc [55, 40]. a general relativistic description of the ksze has been given in the framework of the general boltzmann equation [29] and in the relativistic covariant formalism [38]. sz effect polarization. the ics process yields naturally a polarized upscattered radiation field. the polarization π of the sze arises from various dynamical 57 sergio colafrancesco and plasma effects [54, 22, 32, 1]: galaxy clusters transverse motion (πk ∝ β2t τ in the rayleigh-jeans, rj, regime), transverse motions of plasma within the cluster (πv ∝ βtτ2 in the rj regime) and multiple scattering between electrons and cmb photons within the cluster (πth ∝ θτ2 in the rj regime for the thermal sze with θ = kte/mec 2). a general, covariant, relativistic derivation of the sze polarization for thermal, non-thermal and relativistic plasma can be derived [22] and generalizes the non-relativistic derivation [54] in a way similar to the general derivation of the sze [6] previously discussed. 3 astrophysical and cosmological impact studying the sze in the atmospheres of various cosmic provides many insights on their energetics, pressure and dynamical structure. the combination of sze with other emission mechanisms related to the same particle distribution (i.e., synchrotron, high-e ics and bremsstrahlung emission) provides further information on the radiation, matter and magnetic fields that are co-spatial with the electrons producing the sze. these properties of the sze concern various cosmic structures, from galaxy clusters to radiogalaxy lobes, from galaxy halos to supercluster. the redshift-independent nature of the sze allows to use it as a powerful cosmological probe yielding constraints on the evolution of cluster abundance, cosmological parameters, the dark energy equation of state, the homogeneity of the universe, the properties of cosmological magnetic fields [14]. observations aimed at these challenges must however exploit the whole physical information contained in the spectral and spatial features of the sze. 3.1 simple sze astrophysics the sze has been searched in galaxy clusters since it was originally proposed [55, 56] using various techniques (see [14] for a recent review). ground-based sze experiments (e.g., spt, act, apex, ami, gbt, among others) provided excellent results in terms of imaging and blind search surveys with their low frequency, multiple-band observations, but they do not have neither true spectroscopic capabilities nor a wide frequency band, and they are not sensitive to the high-ν range ( ∼> 400 ghz) of the sze signals, which is crucial to exploit the astrophysical information contained in the sze (see fig. 1, [14]). ground-based instruments widely improved the source statistics (crucial to obtain cosmological information using the sze) and the angular resolution of sze images (crucial to disentangle the extended sze signal from point-source contamination), but add little to the physical specification of the detected sze sources, and therefore they need x-ray and optical follow-up to fully characterize the physical parameters derived from sze observations. figure 1: thermal sze spectrum of galaxy clusters with different plasma temperatures (as indicated) and the same value of the optical depth. the low-ν part (ν ∼< 220 ghz) of the spectrum depends mostly on the total compton parameter y ∝ ∫ d`ptot with no strong spectral dependence on the temperature. the high-ν part of the spectrum (at ∼> 300 ghz) shows a strong spectral dependence from the plasma temperature [17]. typical frequency bands where the sze is observed from the ground are shown as blue-cyan bands, while the region accessible from space observations is shown by the gray shaded area. 3.2 the road to astrophysics the planck satellite allowed for the first time to access a wide frequency range in the study of the sze. planck early sze science observations yielded 189 sze sources with s/n > 6 which provides the first sze measure for ∼ 80 % of the known galaxy clusters, and 20 additional new clusters (see arnaud at this meeting). this is the largest sample so far of sze detected clusters. planck detected sze clusters are followed-up with a multi-frequency observation program in the x-rays, sze, optical bands to obtain confirmation, redshift estimation and estimates of the global physical parameters. these results show that the sze selection is a very powerful method for the detection of new distant and very massive clusters. planck also unveiled a population of dynamically perturbed clusters at z ∼> 0.3, 58 multi-frequency study of the sz effect in cosmic structures possibly underrepresented in x-ray surveys. the information collected so far strengthen our overall view of the icm properties and mass content in galaxy clusters. most of these results are discussed in the early and intermediate papers [41, 42, 43, 45, 46, 47] and more analyses are still coming. the use of planck sze results also allowed to extend the sample of radio halo clusters with combined radio, sze and x-ray data [23]. the correlation ysz−lx probes the existence of a substantial amount of non-thermal pressure in clusters that also requires a correlation x ∝ l−0.96x between the nonthermal to thermal pressure ratio x = pnon−th/pth and the cluster x-ray luminosity lx [23]. the nature of the non-thermal electron population could be only probed with additional high-ν spectroscopic studies. the herschel satellite (co-eval with planck) has been able to observe the sze in a few pointed clusters with the spire instrument equipped with an fts spectrometer working in the frequency range ∼ 600 − 1200 ghz. the possibility to have sensitive spectroscopic measurements in these high-frequency bands opens the way to the deep astrophysical exploitation of the sze. as an example, the additional data points on the sze spectrum of the bullet cluster observed with herschelspire [60] allowed to establish a number of properties for the thermal and non-thermal plasma superposition in the atmosphere of this strong merging cluster (see, e.g. [18, 48, 49]. a consequence of such superposition is that cluster temperature distribution has a high value of of the temperature standard deviation σt , as in the case of the bullet cluster where it is found σt = 10.6 ± 3.8 kev [50]. this result shows that the temperature distribution in the bullet cluster is strongly inhomogeneous along the line of sight and provides a new method for studying galaxy clusters in depth. these studies have been possible because the access to the very high-ν part of the sze spectrum contains detailed information on the relativistic effect on the single thermal plasma and on the presence of additional plasmas of either thermal or non-thermal nature (see discussion in [18, 21]. planck and herschel observations of the sze opened a rich field of investigation that will fully blossom in the next years with the full exploitation of spatiallyresolved spectroscopic sze observations. in the following we discuss some of the astrophysical and cosmological studies that will be possible with a spectropolarimetric study of the sze. 4 astrophysical impact studying the sze in various cosmic atmospheres provides many insights on their energetics, pressure and dynamical structure. the combination of sze with other emission mechanisms related to the same particle distribution (i.e., synchrotron, high-e ics emission, bremsstrahlung emission) provides further information on the radiation, matter and magnetic fields that are co-spatial with the electrons producing the sze in various cosmic structures. 4.0.1 galaxy clusters precise observations of the sze at mm and sub-mm wavelengths are crucial for unveiling the detailed structure of cluster atmospheres, their temperature distribution, and the possible presence of suprathermal and/or nonthermal plasma because the high-frequency part (i.e. at ν ∼> 350 ghz or x ∼> 6) of the sze spectrum is more sensitive to the relativistic effects of the electron momentum distribution [6, 10, 17]. this is even more so for galaxy clusters with a complex plasma distribution as found for powerful merging clusters, like the exemplary case offered by the bullet cluster (1es0657-56) [18]. powerful merging events in galaxy clusters can, in fact, produce an additional high-t plasma distribution (if the electron acceleration time scale at the merging shocks is longer than their equilibration time scale [57]), or an additional nonthermal population (produced either in a merging process with a very short acceleration time scale or by secondary electrons produced by p-p collisions, after the high-e protons have been accelerated by the merging and accumulate in the cluster region on long time scales [58]). the quasi-stationary case provided by the competition between particle thermalization and stochastic acceleration and momentum diffusion [26] can develop a subrelativistic electron distribution tail and can produce suprathermal regions in the cluster atmosphere. a quantitative estimate of the temperature inhomogeneity (stratification) along the line of sight is possible using sze data only providing a measure the temperature standard deviation of the cluster plasma along the line of sight. we found that the bullet cluster has a temperature standard deviation of 10.6±3.8 kev [50]. this result (obtained for the first time with sze measurements) shows that the temperature distribution in the bullet cluster is strongly inhomogeneous and provides a new method for studying galaxy clusters in depth. study of the multifrequency (from ∼ 30 to ∼ 850 ghz) sze signal observed in the bullet cluster shows, in fact, the presence of a thermal plasma at ∼ 13.9 kev coexisting with a second plasma component, either at higher temperature (≈ 25 kev) or, more preferably, of a nonthermal origin [18] (see fig.2). additional observations of the bullet cluster at ν ∼ 400 ghz with a precision ∼< 1% of the expected signal will be able to further distinguish between the two cases of non-thermal powerlaw or suprathermal tail [18]. sze observations over a wide frequency range, and especially with high sensitivity in the high-ν range, can 59 sergio colafrancesco also add relevant information on the electron distribution function (df) in the icm, a subject that even though relevant for a proper analysis of the sze has not been addressed in details so far. the relativistic kinetic theory, on which the df derivation is based, is still a subject of numerous debates (see discussion in [48]). sze observations can separate the sze spectrum caused by a departure from the diffusive approximation based on the kompaneets approach [31] from those which are due to using a relativistic correct df instead of a maxwell-boltzman df (see fig.3) and therefore set constraints to the actual electron df [48]. this analysis is best fulfilled in hot massive clusters because the sze intensity change due to using a relativistic correct df instead of a maxwell-boltzman df are much larger in hot clusters due to the fact that relativistic sze corrections scale as ∝ t5/2. a method used to derive the df of electrons using sze multi-frequency observations of massive hot clusters [48] makes use of fourier analysis representation of the approximate electron df whose parameters are best fitted using observations in the (optimal) frequency channels at 375, 600, 700, 857 ghz. figure 2: the sze spectrum at the bullet cluster center modeled with a thermal plus nonthermal plasma: thermal plasma with kt = 13.9 kev and τ = 1.1×10−2 (dot-dashed); nonthermal plasma with p1 = 1, s = 2.7 and τ = 2.3 × 10−4 (dotted); total sze produced by the sum of the two plasmas (solid). a morphological analysis of the sze observed at various frequencies adds relevant information to assess the pressure and energy density structure of cluster atmospheres. morphological sze differences are particularly evident for clusters undergoing violent mergers that create large inhomogeneities of the electron df. sze intensity maps of merging clusters obtained from hydrodynamical simulations show that the morphology of the sze intensity maps observable with laboca (at 345 ghz) and herschel-spire (at 857 ghz) are rather different [49] (see fig.4). for a bullet-like cluster, the sze intensity map at 857 ghz has a spatial feature caused by the presence of the cold bullet-like substructure seen also in the x-ray surface brightness map. however, this cold substructure is not present in the sze intensity map at 345 ghz. this is a consequence of the relativistic effects of the sze and shows that observations of the sze intensity maps at very high frequencies can reveal complex pressure substructures within the atmospheres of massive galaxy clusters. this result shows that the analysis of the sze signal at 857 ghz, correlated with lower-ν observations offers a promising method for unveiling high-t regions in massive merging clusters using available experiments like, e.g laboca and herschel-spire. figure 3: the sze intensity spectra for a massive cluster with a temperature of kte = 15.3 kev for juttner (solid) and maxwell-boltzman (dashed) dfs. the non-relativistic sze spectrum (solid) is also shown for comparison. figure from [48] in more relaxed clusters spectroscopic measurements of the sze over a wide frequency band allow to derive precise information on the temperature distribution and on the cool-core nature independently of x-ray priors [17] and hence reconstruct the full set of cluster physical parameters [25]. spectro-polarimetry measurement of the sze are able to add further information on the transverse plasma motions within the cluster and on the pressure substructure of the plasma. sze polarization signals in galaxy clusters are quite low and typically below mjy (or ≈ µk) level even for high-t clusters [14] but it is interesting to note that the sze polarization in cluster has quite different spectra w.r.t. the intensity sze spectrum, a property that requires to use spectropolarimetry to fully disentangle the physical information contained in the sze. combining intensity and polarization observations of the sze can uncover unique 60 multi-frequency study of the sz effect in cosmic structures details of the 3d (projected and along the line of sight) velocity structure of the icm, of its 3d pressure structure and of the influence of a structured magnetic field in the stratification of the icm, and therefore provides a full tomography of cluster atmospheres. analogously, the combination of the intensity and polarization observations of the kinematic sze (and its frequency dependence) can yield crucial information on the 3d distribution of the cosmological velocity field traced by galaxy clusters. specifically, the ratio ∆ith/πth yields direct information on the plasma optical depth τ, and the ratio ∆ith/πv on the combination τ ·βt, thus allowing to use intensity and polarization sze measurements to fully disentangle the pressure and velocity structure of the cluster atmospheres. sze polarization measurements are quite difficult to obtain with present-day experiments and they are also at the limit of next generation experiments. however, stacking analysis of even small samples (order of ∼ 20) of hot and dense galaxy clusters would allow to determine statistically the polarization signals of the thermal sze for clusters with kt > 10 kev and τ > 0.03 in the optimal frequency range ≈ 90 − 250 ghz. figure 4: from top to bottom. the sze signal to noise ratio map for the cluster 1e0657-558 at 345 ghz smoothed to the resolution of laboca and at 857 ghz smoothed to the resolution of herschel-spire. figures from [49]. 4.0.2 cluster cavities the atmospheres of galaxy clusters often show the presence of bubbles filled with high-e particles and magnetic field that are sites of bright radio emission and produce cavities in their x-ray emission distribution. cavities with diameters ranging from a few to a few hundreds of kpc have been observed by chandra in the x-ray emission maps of several galaxy clusters and groups [3, 34]. while the properties of these cavities and of the relativistic plasma they contain is usually studied by combining x-ray and radio observations, an alternative and efficient strategy is to study the consequences of the sze produced by the high-energy electrons filling the cavities [7, 39] whose amplitude, spectral and spatial features depend on the overall pressure and energetics of the relativistic plasma in the cavities. as an example, the overall sze observable along the line of sight (los) through a cluster containing cavities (see fig.5 for the case of the cluster ms0735.6+7421) is the combination of the non-thermal sze produced by the cavity and of the thermal sze produced by the surrounding icm. due to the different ν-dependence of the thermal and non-thermal sze, the non-thermal sze from a cluster cavity shows up uncontaminated at frequencies ν ≈ 220 ghz: at this frequency, in fact, the overall sze from the cluster reveals only the ics of the electrons residing in the cavities without the presence of the intense thermal sze dominating at lower and higher frequencies. the cavity’s sze becomes dominant again at very high-ν (x ∼> 14 or ν ∼> 800 ghz) where the nonthermal electrons dominate the overall ics emission (see fig.5). the cavity’s sze is more spatially concentrated than the overall cluster sze because it is only emerging from the cavity regions: this fact allows to study the overall energetics and pressure structure of the cavity’s high-e particle population and the b-field structure in combination with x-rays and radio images. the observation of the crossover of the non-thermal sze from the cavities (which depends on the value of emin(p1) or, equivalently, on the value pcavity) provides a way to determine the total pressure and hence the nature of the electron population within the cavity [7], an evidence which adds crucial, complementary information to the x-ray and radio analysis. if cluster cavities contain a high-t (∼ 109 − 1010 k) plasma, the sze flux from cocoons in the central part of a distant elliptical and a nearby galaxy cluster are of the same order. for a high-t plasma, the cocoon’s sze spectrum is rather flat at high-ν resembling the shape of the non-thermal sze from cavities. in this high-t plasma model, however, no strong radio emission at ν ∼> 1 ghz (as instead observed) is expected from the cocoon, unless the cocoon’s b-field is very high b ∼> 103µg. 4.0.3 radiogalaxy lobes studies of (giant) radio-galaxy (rg) lobes (see, e.g., [28, 30, 24, 4]) have shown that these extended structures contain relativistic electrons that are currently available to produce both low-ν synchrotron radio emission and ics of the cmb (as well as other radiation background) photons. as a consequence, an sze from the lobes of rgs is inevitably expected [11]. such nonthermal, relativistic sze has a specific spectral shape that depends on the shape and energy extent of the spectrum of the electrons residing in rg lobes. the 61 sergio colafrancesco sze emission from rg lobes is expected to be co-spatial with the relative ics x-ray emission [11] and its spectral properties are related to those of the relative ics x-ray emission. in fact, the spectral slope of the ics x-ray emission αx = (α− 1)/2 (where fics ∼ e−αx ) can be used to set the electron energy spectral slope α (where ne ∼ e−α) necessary to compute the sze spectrum, and to check its consistency with the synchrotron radio spectral index αr = (α − 1)/2 (where fsynch. ∼ e−αr ), that is expected to have the same value [11, 19]. figure 5: top. the geometry of the cavities in the cluster ms0735.6+7421. bottom. the sze spectrum has been computed at a projected radius of ≈ 125 kpc from the cluster center where the los passes through the center of northern cavity. the thermal sze (blue), the non-thermal sze from the cavity (black) and the total sze (red) are shown. the non-thermal sze is normalized to the cavity pressure p = 6 · 10−11 erg cm−3, and is shown for various values of p1. figure from [7] the sze in rg lobes has not been detected yet: only loose upper limits have been so far derived on the sze from these sources (see [2, 59]). a detection of the sze from rg lobes can provide a determination of the total energy density and pressure of the electron population in the lobes [11] allowing to determine the value of emin once the slope of the electron spectrum is determined from radio and/or x-ray observations. sze measurements provide a much more accurate estimates of the electron pressure/energy density than with other technique like ics x-ray emission or synchrotron radio emission, since the former can only provide an estimate of the electron energetics in the high-energy part of the electron spectrum, and the latter is sensitive to the degenerate combination of the electron spectrum and of the magnetic field in the radio lobes. the combination of sze observations (that depend on the electron distribution and on the known cmb radiation field) and the radio observations (which depend on the combination of the electron distribution and of the magnetic field distribution) provides an unbiased estimate of the overall b-field in the lobe by using the ratio fradio/fsze ≈ eb/ecmb, that is more reliable than that obtained from the combination of ics x-ray (or gamma-ray) and radio emission [19]. the spatially resolved study of the sze and synchrotron emission in rg lobes also provide indication on the radial behaviour of both the leptonic pressure and of the magnetic field from the inner parts to the boundaries of the lobes. study of the pressure evolution in rg lobes can provide crucial indications on the transition from radio lobe environments to the atmospheres of giant cavities observed in cluster atmospheres, which seem naturally related to the penetration of rg jets/lobes into the icm (see fig.5). a substantial sze polarization is also expected in rg lobes due to both coherent transverse motions of the plasma along the jet/lobe direction and to the electron pressure substructures induced by e.g. plasma inhomogeneities and magnetic field turbulence. the transverse velocity-induced polarization is πv ∝ τrel(βtτrel), and the multiple scattering induced polarization is πτ ∝ τrelprel where prel is the pressure of the relativistic electron distribution. observations of the sze and its polarization in rg lobes can yield, therefore, direct information on electrons τrel and βt in the rg lobe. 4.0.4 galaxies hot gas trapped in a dm halo can produce a sze. a typical galaxy halo might hence show an integrated thermal sze at the level of ∼< 0.5 mjy arcmin−2 from a plasma with t ∼ 106 k and density ne ∼ 10−3cm−3 extended for ∼ 50 kpc in the inter-stellar medium. a measurement of galaxy halo sze would provide direct information on the mass, spatial distribution and thermodynamic state of the plasma in a low-mass galactic halo, and could place important constraints on current models of galaxy formation. detecting such an extended, low-amplitude signal will be challenging, but possible with sensitive all-sky sze maps. an sze is also expected from galaxy outflows swept by galaxy (hyper-)winds. a thermal sze is expected to arise from the shocked bubble plasma in a galaxy wind described by a simple, spherical blast wave model [52]. 62 multi-frequency study of the sz effect in cosmic structures however, such simple recipe for the sze from galaxy winds is to be modified by the presence of cosmic rays and magnetic field in the expanding wind leading on one side to a more complex sze spectrum, and on the other side to an amplification of the overall sze at high frequencies, thus increasing the detection probability. sze observations from galaxy winds will be possible with high-sensitivity and high-resolution telescopes like alma and ska.     figure 6: top. thermal sze spectrum (continuous line in bottom panel), compared to the atmospheric transmission of a dry, cold, atmosphere (top panel, pwv= 0.5 mm), and to spectra of the nonthermal sze (dotted line), of cmb anisotropy and kinematic sze (dashed line), and of dust anisotropy (dotdashed line). the parameters of the different spectra are for τth = 5 × 10−3, te = 8.5 kev, ∆tcmb = 22µk, τnon−th = 10 −4, α = −2.7, p1 = 1.4 mev/c, ∆id(150ghz) = 600 jy. the frequency coverage of different experiments [25] is shown as dotted horizontal lines, labelled with the experimental configuration number. bottom. simulated data sets for the spectroscopic configurations ec5 (differential cold fts on a l2-orbit satellite, with cold telescope). 5 cosmology and fundamental physics the redshift-independent nature of the sze allow to use this effect as a powerful cosmological probe [14] by using both the redshift evolution of cluster abundance and direct probes of cosmological parameters. the recent planck results however indicate a tension at ∼ 3σ level on the value of σ8 vs. ωm comparing sze counts and cmb results that might be mitigated by other cosmological effects like, e.g., a more complex cluster bias dependence, the existence of massive neutrinos with ∑ mν ∼ 0.2 ev (see rubino-martin in these proceedings) or by addressing more in details the complex astrophysics issues still pertaining cluster formation and evolution. this last possibility requires to have new high-quality sze cluster samples studied with full spectra-polarimetric capabilities and in the high-ν range of the sze spectra where the astrophysical effects are more evident. fig.6 shows the frequency coverage requirements for a full spectroscopic sze mission and the outcomes of a simulated thermal sze spectrum from a kte = 8.5 kev cluster [25]. the sensitivity and spectro-polarimetry power of such an experimental configuration (ec5) will allow to address fundamental questions in cosmology and fundamental astrophysics with the sze, in addition to standard cosmological probes [14]. a wide discussion on the cosmological relevance of the sze for probing the concordance cosmological scenario, the search for the nature of dm, cosmological magnetic fields and cosmological velocity fields can be found in [14]. here we discuss the additional application of the spectro-polarimetric observations of the sze to measure the cmb multiplies and test the assumption of the homogeneity of the universe. high-sensitivity observations of the (thermal and kinematic) sze and its polarization can be used to test the homogeneity of our universe through probes of the copernican principle (cp) [33, 23]. in fact, large variations of the thermal sze induced by cmb photons that have temperature significantly different from the blackbody temperature we observe directly, and that arrive at galaxy cluster from points inside our light cone, could indicate a violation of the cp and homogeneity. analogously, large variations of the cmb dipole measured by the cluster ksze could also indicate a violation of the cp and homogeneity. the sze polarization contains more refined information on the cmb temperature and thus it can also provide a powerful probe of cp and homogeneity. thus, observations of large sze and of its polarization w.r.t. to the expectations of the sze produced from a pure blackbody cmb spectrum might provide indications of a non-flrw universe. the combination of an ec5 experiment with radio observations of the low-ν part of the sze spectrum can also be used to probe even more fundamental aspects of astrophysics, like the fundamental properties of the photon. measurements of the sze with sensitivity of order of < 0.1 jy, in the range ∼ 10 − 50 ghz, can set very stringent constraints on the photon decay time [20]. this frequency band is also the one less affected by other sources of astrophysical contamination, and it will be best explored with the advent of the high63 sergio colafrancesco sensitivity ska telescope. at higher frequencies (i.e., ∼ 120 − 180 and ∼ 200 − 300 ghz), there are other spectral windows where the sze method is again competitive if not advantageous in this respect. the necessary sensitivity in the high-ν range can be achieved with the next coming millimetron space mission [14], whose ec5-like configuration is also advantageous in disentangling other astrophysical sources of contamination (e.g., sources of non-thermal sze and/or special distortions of the sze due to multiple-temperature regions) that could contaminate these measurements. figure 7: difference between the standard thermal sze calculated for a2163 in 2–45 ghz range and integrated in a 1 arcmin2 area and the one modified for values e∗ = 5×10−9 ev (solid line), 3×10−9 ev (dash 3 dots), 2 × 10−9 ev (dot-dashed), and 1 × 10−9 ev (dashed). the sensitivity achievable with ska in ∼ 30 hours (red thick solid line) and in ∼ 260 hours (blue thick long-dashed line) are shown [20] 6 future directions the impact of sze observation for astrophysical and cosmological application is steadily increasing since the advent of dedicated experiments and large-scale surveys in the mm and sub-mm frequency range. the quality and the spatial resolution of sze images is reaching the level of arcsecs and the spectral coverage in systematically extending in the sub-mm region where highest spatial resolutions can be achieved with a wide spectral coverage able to decipher the physical details of the electron distribution in the atmosphere of cosmic structures. new paths of theoretical investigations are underway and concern both the detailed study of the ics mechanism and the impact of the various plasma structure and fields in cosmic structures. the possibility to perform precise measurements of the various sze signals and to extract the relevant astrophysical information depends crucially on the capability to have spatially-resolved spectral (and polarization) observations of sze sources over a wide ν band, from radio to sub-mm. in particular, the important condition for such study is to have a wide-band continuum spectroscopy (polarimetry) and especially a good spectral coverage and sensitivity in the high-ν band, where most of the astrophysical effects reveals more clearly. spatially resolved spectroscopic and polarimetric observations of the sze in the frequency range from ∼ 100 ghz to ∼ 1 thz, complemented by analogous observations in the radio band ∼ 1 − 30 ghz, are the key to improve our understanding of the structure of cosmic atmospheres through analysis of the intensity and polarized sze signal, and will allow to use this technique to probe the fundamental properties of the universe. acknowledgement s.c. acknowledges support by the south african research chairs 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http://dx.doi.org/10.1086/498227 http://dx.doi.org/10.1086/591723 introduction the physics of the sz effect astrophysical and cosmological impact simple sze astrophysics the road to astrophysics astrophysical impact galaxy clusters cluster cavities radiogalaxy lobes galaxies cosmology and fundamental physics future directions frascati workshop 2013 the tenth international workshop on multifrequency behaviour of high energy cosmic sources palermo, sicily, italy, may 27 – june 1, 2013 edited by franco giovannelli & lola sabau-graziati workshop on multifrequency behaviour of high energy cosmic sources palermo, sicily, italy may 27 june 1, 2013 organizing institutions istituto di astrofisica e planetologia spaziali iaps-inaf, roma, italy dpt de cargas utiles y ciencias del espacio, dcuce-inta madrid, spain e.o. hulburt center for space research, hca-nrl, washington d.c., usa max-planck institut fr extraterrestrische physik, mpe, garching, germany st. john’s college, sjc, annapolis md, usa institute für astronomie und astrophysik, karls eberhard, universität sand 1, tübingen, germany ascr astronomical institute, ondřejov, czech republic ctu czech technical university, prague, czech republic scientific organizing committee james howarth beall, hca-nrl & sjc franco giovannelli, inaf-iaps (chairperson) thomas boller, mpe lola sabau-graziati, inta-dcuce andrea santangelo, iaa, universitt tübingen rené hudec, ascr & ctu, prague, czech republic local organizing committee franco giovannelli, inaf-iaps (istituto di astrofisica e planetologia spaziali) (chairperson) daniela giovannini, cnr-ibcn (istituto di biologia cellulare e neurobiologia) paolo persi, inaf-iaps (istituto di astrofisica e planetologia spaziali) francesco reale, cnrisc (istituto dei sistemi complessi) ii foreword the workshop on multifrequency behaviour of high energy cosmic sources is a classical biennial meeting held in the same historical place for our workshops: the beautiful and wild island of vulcano. unfortunately, because of the eolian hotel is closed for repairs, this tenth edition of our workshop has been developed in palermo at the splendid hotel la torre. we are completely convinced that a deep discussion on the multifrequency behaviour of high energy cosmic sources, defined mostly by the physical processes rather than the types of objects, is the most powerful way for a better and faster development of our knowledge of the physics governing our universe. for this purpose, we invited distinguished colleagues representative the most important space– and ground–based experiments of the present generation. they painted the whole panorama on the experimental results and on their interpretation by using current models and/or developing new ideas for a better comprehension of the detected phenomena. during the discussion of a lot of new results, obtained along most of the electromagnetic spectrum, we have learned that the astrophysics of galactic sources is still extremely interesting and unavoidable. indeed it provides crucial information on the same physical processes occurring also in the extra– galactic sources, with obvious different scale factors. moreover, new results coming from very high energy γ-ray experiments have shown the richness — and its potential increase — of the sky at those energies having detected hundreds cosmic sources until few years ago completely unknown, but hopefully to be detected since the cosmic ray spectrum extends up to energies as high as ≥ 1020 ev. the fluxes of very high energy γ-rays are attenuated because of their interactions with the cosmic radio, microwave, infrared and optical radiation fields. measurements of the flux attenuation can then provide important information on the distribution of such fields. then, very high energy γ-ray astronomy is the most powerful tool for solving the problem about the origin of cosmic rays. many thanks to the authors, that accepted our reminders recalled at the beginning of the workshop, for keeping to schedule. however, in spite of the accurate respect of the dead line of a number of authors, we still suffer a delay in publishing these proceedings because of the laziness of a considerable part of other authors. in spite of this, these proceedings will appear roughly within one year from the workshop date. this is the tenth edition of the workshop (tenth if we consider the first historical one, held in vulcano in 1984), whose proceedings are published in the international czech journal ‘acta polytechnica’. we would like to thank all participants and especially the speakers for their active contributions in rendering this workshop updated with their talks, alive with their discussions, and friendly with their attitudes. particular thanks to mr francesco reale for his excellent assistance for informatics in roma, during the preparation of the workshop, and for his faultless assistance in palermo. a special thank to daniela giovannini for her help during the critical moments of the registration of participants. we are grateful to the pianist maèstra valentina usai, to the violinist maèstro alessandro perpich and to the actresses lisa colosimo and flavia giovannelli for their wonderful performance ”hic habitat felicitas: mozarterie from the kingdom behind”. iii many thanks to: the director of the istituto di astrofisica e planetologia spaziali of the istituto nazionale di astrofisica (iaps-inaf) (dr pietro ubertini); the caen s.p.a. (viareggio) who partially supported the organization of this workshop. particular thanks to the plan nacional de i+d+i of spain for partially supporting the publication of this book. finally, we would like to thank mr daniele inzerillo, the chef of the splendid hotel la torre, who delighted us with his creations during our stay and to the whole staff of the hotel for their professionalism and kindness. the editors franco giovannelli & lola sabau-graziati iv list of participants amati lorenzo , inaf–iasf, bologna, italy amati@iasfbo.inaf.it ambrosino filippo , la sapeinza university, roma, italy filippo.ambrosino@roma1.infn.it anderson gemma , university of southampton, uk gemma.anderson@hotmail.com arnaud monique , cea service d’astrophysique, gif sur yvette, saclay, france monique.arnaud@cea.fr aschenbach bernd , prv vaterstetten, germany bernd.aschenbach@t-online.de auriemma giulio , universit degli studi della basilicata, italy giulio.auriemma@cern.ch barret didier , irap, toulouse, france didier.barret@irap.omp.eu barrière nicolas , space sciences lab uc berkeley, ca, usa barriere@ssl.berkeley.edu beall james h. , ssd-nrl, washington dc, usa beall@sjca.edu behar ehud , technion. haifa, israel behar@physics.technion.ac.il beilicke matthias , washington university in st.louis, la, usa beilicke@physics.wustl.edu bisnovatyi-kogan gennady , space research institute, iki, moscow, russia gkogan@iki.rssi.ru v blay pere , gace–lpi–universidad de valencia, spain pere.blay@uv.es boller thomas , mpe, garching, germany bol@mpe.mpg.de bordas pol , iaat & isdc, tübingen, germany pol.bordas@uni-tuebingen.de buckley david , southern african large telescope, cape town, south africa dibnob@saao.ac.za castellina antonella , inaf–oato and infn, torino, italy castelli@to.infn.it castro cerón josé maŕıa , villanueva de la cañada, madrid, spain josemari@alumni.nd.edu colafrancesco sergio , school of physics wits university, johannesburg, south africa sergio.colafrancesco@wits.ac.za costamante luigi , dipartimento di fisica, università di perugia, italy luigic2011@gmail.com dar arnon , technion, israel institute of technology, haifa, israel arnon@physics.technion.ac.il de laurentis mariafelicia , dpt scienze fisiche, università ”federico ii”, napoli, italy felicia@na.infn.it de ugarte postigo antonio , iaa–csic, granada, spain adeugarte@gmai.com della valle massimo , inaf–osservatorio astronomico di capodimonte, napoli, italy dellavalle@na.astro.it di sciascio giuseppe , infn–sezione di roma tor vergata, italy disciascio@roma2.infn.it vi ebisawa ken , jaxa/isas, sagamihara, japan ebisawa@isas.jaxa.jp moshe elitzur , university of kentucky, ky, usa moshe@pa.uky.edu fargion daniele , dipartimento di fisica, università ”la sapienza” di roma, italy daniele.fargion@roma1.infn,it federici memmo , inaf–iaps, roma, italy memmo.federici@iaps.inaf.it georganopoulos markos , university of maryland, baltimore county, nasa–gsfc, baltimore, md, usa georgano@umbc.edu giovannelli franco , inaf–iaps, roma, italy franco.giovannelli@iaps.inaf.it gómez de castro ana inés , universidad complutense de madrid, spain aig@mat.ucm.es guainazzi matteo , european space agency, villanueva de la cañada, spain matteo.guainazzi@sciops.esa.int harms benjamin , university of alabama, al, usa bharms@bama.ua.edu hudec rené , astronomical institute academy of science, ondřejov & ctu prague, czech republic rene.hudec@gmail.com iglesias-groth susana , instituto de astrof́ısica de canarias, la laguna, spain sigroth@iac.es katsuda satoru , riken, wako, saitama, japan satoru.katsuda@riken.jp vii kreykenbohm ingo , dr karl remeis observatory & ecap, bamberg, germany ingo.kreykenbohm@sternwarte.uni-erlangen.de krumpe mirko , eso, garching, germany mkrumpe@ucsd.edu kundt wolfgang , argelander institute of bonn university, germany wkundt@astro.uni-bonn.de lott benoit , cenbg, gradignan, france lott@cenbg.in2p3.fr maeda keiichi , kavli ipmu, university of tokyo, japan keiichi.maeda@ipmu.jp marshall herman , mit, cambridge, ma, usa hermanm@space.mit.edu martino bruno luigi , cnr–iasi (istituto analisi dei sistemi ed informatica), roma, italy brunolmartino@gmail.com meintjes pieter , university of the free state, bloemfontein, south africa meintjpj@ufs.ac.za meli athina , university of gent, dept. of physics and astronomy, gent, belgium ameli@phys.uoa.gr merafina marco , dipartimento di fisica università ”la sapienza” di roma, italy marco.merafina@roma1.infn.it milano leopoldo , dipartimento di scienze fisiche, università ”federico ii”, napoli, italy milano@na.infn.it morihana kumiko , institute of physical and chemical reserch (riken), wako, saitama, japan morihana@crab.riken.jp viii morselli aldo , infn roma tor vergata, roma, italy aldo.morselli@roma2.infn.it motch christian , cnrs–observatoire astronomique de strasbourg, france christian.motch@unistra.fr muñoz-tuñon casiana , instituto de astrof́ısica de canarias, la laguna, spain cmt@iac.es nebot ada , observatoire astronomique de strasbourg, france ada.nebot@astro.unistra.fr nozawa takaya , kavli ipmu, university of tokyo, kashiwa, japan takaya.nozawa@ipmu.jp orlando salvatore , inaf–osservatorio astronomico di palermo, italy orlando@astropa.inaf.it panagia nino , space telescope science institute, baltimore, maryland, usa panagia@stsci.edu parisi pietro , inaf–iaps, roma, italy pietro.parisi@iaps.inaf.it persi paolo , inaf–iaps, roma, italy paolo.persi@iaps.inaf.it perucho manel , dpt d’astronomı́a i astrof́ısica, universitat de valència, spain manel.perucho@uv.es piro luigi , inaf–iaps, roma, italy luigi.piro@iaps.inaf.it pittori carlotta , inaf–oar and asdc, frascati, italy pittori@asdc.asi.it reale francesco , cnr–isc, roma, italy francesco.reale@sic.rm.cnr.it ix rebolo rafael , instituto de astrof́ısica de canarias, la laguna, spain rrl@iac.es regis marco , università di torino and infn–torino, italy regis.mrc@gmail.com reimer anita , universität innsbruck, austria anita.reimer@uibk.ac.at ribó marc , universitat de barcelona, spain mribo@am.ub.es rodŕıguez fŕıas , space and astroparticle group, uah–madrid, alcal de henares, spain dolores.frias@gmail.com rosinska dorota , kepler institute of astronomy, university of zielona gora, poland dorota@astro.ia.uz.zgora.pl rubino-martin josé alberto , instituto de astrof́ısica de canarias, la laguna, spain jalberto@iac.es sabau-graziati maŕıa dolores , inta–división de cargas utiles y ciencias del espacio, torrejón de ardoz, madrid, spain sabaumd@inta.es sanchez-portal miguel , herschel science centre, esac/european space agency, villanueva de la cañada, spain miguel.sanchez@sciops.esa.int sani eleonora , inaf-oa arcetri, firenze, italy sani@arcetri.astro.it santangelo andrea , institute of astronomy and astrophysics, kepler center, university of tübingen, germany andrea.santangelo@uni-tuebingen.de sasaki manami , institute of astronomy and astrophysics, kepler center, university of tübingen, germany sasaki@astro.uni-tuebingen.de x schartel norbert , esa/xmm–newton soc, villanueva de la cañada, madrid, spain norbert.schartel@sciops.esa.int sikora marek , pan–nicolaus copernicus astronomical center, warszawa, poland sikora@camk.edu.pl šimon vojtěch , astronomical institute academy of science, ondřejov, czech republic vojtech.simon@gmail.com van soelen brian , university of the free state, bloemfontein, south africa vansoelenb@ufs.ac.za spurio maurizio , university of bologna and infn–bologna, italy spurio@bo.infn.it stanev todor , university of delaware, newark, de, usa stanev@bartol.udel.edu tluczykont martin , university of hamburg, germany tluczykont@gmail.com tornambé amedeo , inaf–roma observatory, monteprozio catone, italy tornambe@oa-teramo.inaf.it troja eleonora , nasa–gsfc/cresst, greenbelt, md, usa eleonora.troja@nasa.gov tsuru takeshi , physics dpt, university of kyoto, japan tsuru@cr.scphys.kyoto-u.ac.jp ubertini pietro , inaf–iaps, roma, italy pietro.ubertini@iaps.inaf.it vitantoni daniele , dipartimento di fisica università ”la sapienza” di roma, italy daniele.vitantoni@roma1.infn.it walter roland , isdc, university of geneva, versoix, switzerland roland.walter@unige.ch xi yoon sung-chul , university of bonn, germany astroyoon@gmail.com zió lkowski janusz , pan–nicolaus copernicus astronomical center, warszawa, poland jz@camk.edu.pl xii 316 acta polytechnica ctu proceedings 1(1): 316–319, 2014 316 doi: 10.14311/app.2014.01.0316 us astronomical photographic data archives: hidden treasures and importance for high-energy astrophysics rené hudec1,2, lukáš hudec2 1astronomical institute, academy of sciences of the czech republic, cz-25165 ondřejov, czech republic 2czech technical university in prague, faculty of electrical engineering, prague, czech republic corresponding author: rene.hudec@asu.cas.cz abstract we report here on an ongoing investigation of us astronomical plate archives and tests of the suitability of transportable scanning devices for in situ digitization of archival astronomical plates, with emphasis on application in high-energy astrophysics. keywords: astronomical data archives: astronomical photographic archives – spectroscopy: low-dispersion spectra – data mining. 1 introduction there are numerous important astronomical plate archives in the usa, including plate collections that are little known to the community and that have been little investigated in the past. within the framework of a czech-us collaborative project (msmt kontakt amvis me09027), we have recently analyzed some of them, obtaining test scans with the use of a portable digitizing device. digitization is a necessary step for an extended evaluation of the plate data using dedicated programs and powerful computers. several recently found large us negative archives are expected to play important role in high-energy astronomy. motivation for including this subject into multi– frequency studies of high energy sources is as follows. numerous he/vhe/uhe sources are also emitters of optical light, many of them are variable. the astronomical plate archives represent the only method how to study the behavior of the objects over very long (100 years or even more) time intervals (hudec, 1999 and 2012). in addition, huge monitoring times allows to detect and to study rare events such as flares. the databases allow to study prominent spectra and/or spectral changes as well (hudec et al., 2012). 2 the plate archives the us astronomical archival plate collections that we recently visited and investigated include those housed in the following 16 institutions: carnegie observatories pasadena, ca, lick observatory, ca yerkes observatory, wi, mt palomar observatory, ca, pari, rosman, nc (which has a collection of plates from many observatories), kpno tucson, az, cfht waimea, hi, ifa manoa, hi, usno flagstaff, az, usno washington, dc, steward observatory tucson, az, nmsu, las cruces, nm, rosemary hill observatory, university of florida, gainesville, fl and leander mccormick observatory, university of virginia, charlottesville, va, smithsonial archives washington, dc, and hazy space center, dulles, va. our estimate is that there are more than 2 million astronomical archival plates and/or negatives in these archives. there are however many us plate collections not included in our study (robbins and osborn, 2009). we performed a quality check and analyzed these plate archives with emphasis on their scientific, historical and cultural value, which we have found to be enormous. in addition to that, collaboration with several german institutes with astronomical plate archives started, with emphasis on digitization and investigation of relevant photographic collections. 3 transportable digitizing device most of the plate archives that we visited have no plate scanners and lack modern instrumentation in general. as our study includes plate digitization, it was necessary to find a solution. since we were going to travel from europe to the us by air, the obvious option was a transportable digitization device based on a digital camera with a high-quality lens and a stable tripod. this solution has the following advantages over other techniques: the device is easily transportable, and offers much faster scanning and higher repeatability than commercial flatbed scanners, because there are no mov316 http://dx.doi.org/10.14311/app.2014.01.0316 us astronomical photographic data archives: hidden treasures and importance for high-energy astrophysics ing scanner parts. the equipment that we used was as follows: camera: 21 mpx canon eos 5d mark ii, lenses: canon ef 24-70 f/2.8 l usm and canon 70200mm f4, a stable tripod, and a fomei lp-310 professional photographic light table. in last project year, a better custom-made light table based on highly homogeneous led illumination was used, and also a further improved camera (39 mpx) and lenses. the recorded images are then corrected for lens image distortions and for other effects, in order to store research-grade digital images. the achieved pixel size depends on the size of the plate, and fine pixel sizes of order of 25–30 microns or less can be typically obtained for small and medium sized plates (up to 16 cm x 16 cm or 18 cm x 24 cm) while for larger plates larger pixel sizes will be obtained (or, alternatively, the large plate may be covered by several shots to achieve smaller pixel sizes). examples of digitized 16 cm x 16 cm plates are illustrated in figs. 1 and 2 with relevant field recognition and star identification. figure 1: example of digitized bamberg southern sky patrol plate processed by astrometry.net tools: field recognition 4 general picture after visiting the us plate collections mentioned above, we offer a (subjective) list of the major problems found in these archives: (i) the list of us plate collections provided by dr. wayne osborn (robbins and osborn, 2009) was found to be incomplete. we have found valuable plate collections with plates from important telescopes that are not listed, e.g. the two hawaii plate collections in manoa (institute of astronomy) and in waimea (canada-france hawaii telescope cfht). some plate archives have been completely hidden, as their home institutes were in some extreme cases not even aware that they have plate stacks. (ii) in numerous collections, only a very rough estimate of the number of plates can be given, as no exact information about the total number of plates, etc., is available. usually, the real number of plates is higher than the previously available estimate. in general, it is very difficult to give the exact number of plates, due to lack of observation logs and inadequate organization of the plate archives. (iii) in many cases, there is no contact person responsible for the plate archive, and it is difficult to make contact. in some places, it is even difficult to get access. this situation very has a serious adverse effect on efforts to exploit these plate collections for scientific purposes. (iv) for some archives, no information is available, not even an approximate number of plates. in many archives, no plate logs are available; they have either been removed or are lost. the only available information is what is written on the plates or on the envelopes (in some cases, there is not even adequate information on the envelope). we guess that there were originally observation logs, but that these were later separated from the plates and archived in a different location. example: carnegie observatories pasadena (nearly 0.5 million plates), where the logs are probably located in the attic above the library, with difficult access. (v) damaged plates in some archives (mostly due to a partly-released or even complete released emulsion layer), probably due to improper storage (or changes in humidity/temperature over time). we point out however that even these plates can be restored using suitable chemical methods and procedures. (vi) lack of electronic records no lists of plates, the only information is on the plates and/or plate envelopes (vii) many of the archives that we visited suffer from inadequate funding, lack of devices, e.g. no scanners. (viii) we have revealed that many plates have been removed from their home: these plates are usually scattered in private homes and offices, or are being kept by observers (often abroad). numerous plates taken at us observatories have been found in european plate archives. nevertheless, we found highly valuable plates almost everywhere, and the quality of the plates (and hence their scientific potential) is mostly high or even very high, in comparison with the plates in european archives. this is true both for direct images and for spectral images (taken with an objective prism). in addition to stellar images, some of the archives that we visited also include extensive collections of solar images (e.g. carnegie observatories in pasadena) and/or planetary images (unique collection in las cruces). the storage conditions were found to vary from archive to archive, from proper temperature and humidity conditions, to less proper conditions. the main degradation sources found in the plate collections were high levels 317 rené hudec, lukáš hudec of humidity and probably also temperature variation, resulting in partial or complete release of the emulsion layer. the scientific use of the plate archives is negatively impacted by poor access to the plates at some places, and also by the fact that the plates have in most cases not yet been cataloged. figure 2: example of digitized bamberg southern sky patrol plate processed by astrometry.net tools: star identification/astrometry 5 two recently found large us archives with possible major impact on hea the two following databases found within our project are expected to play an especially important role in high-energy astronomy. baker super schmidt camera domed films with 55 degrees diameter fov, limiting magnitude up to 15, dense sampling 20 min, ∼ 110 000 negatives (fig. 5), ∼ 10 years coverage. the essential part of the archive is located at pari (pisgah astronomical research institute) in nc. baker-nunn camera networks negatives fov 30 x 5 degrees, limiting magnitude up to 16, very dense sampling ∼ few sec, more than 1 mil negatives, ∼20 years coverage (figs. 3 and 4). the part of the archive (21 boxes with ∼ 2000 film rolls) is located at smithsonian institution archive in washington, d.c. these data are suitable for wide-field studies, optical transient (ot) searches, and fast variability studies. there are up to 100 000 stars per 1 full frame. these data represent direct images suitable for photometry. however, one baker nunn kwajalein slit-less spectrograph camera for u.s. air force was designed to provide unique low–dispersive spectra. this baker nunn missal reentry tracking camera consisted of a three-element 20-inch entrance aperture corrector lens system. mounted on top of the corrector lens assembly was a multi-set of triangular prisms converting the incoming sky image into a spectrum image. figure 3: dismantling of the usaf baker–nunn camera at the evergreen aviation & space museum, mcminnville, oregon, united states figure 4: example of digitized film from sao baker– nunn camera the corrected spectral image is then projected down to a 31-inch f/1 spherical primary mirror. the focusing converging image was then reflected back up to a 4-inch x 6-inch x 0.6-inch thick mosaic of fiber-optics making film supporting focusing plate. the resulting database is expected to has applications in astronomy as well. 6 summary 16 us astronomical plate archives were visited within the amvis czech-us collaborative project. the quality of the plates and their scientific, historical and cultural value were investigated for possible inclusion in the us astronomical plate repository at pari, nc. some of these archives (e.g. baker super schmidt, baker–nunn, cfht waimea, ifa manoa) were un318 us astronomical photographic data archives: hidden treasures and importance for high-energy astrophysics known to the astronomical community before our study, and the two first mentioned are expected to play important role in high–energy astronomy because of their very wide–field and very fine time resolution. selected plates were digitized using a transportable scanning device. all the archives that we visited have plates that are scientifically valuable, and in many cases unique. the plates are however mostly hidden from the astronomical community, and the plates have not yet been catalogued. the total number of plates is higher than expected in many of the locations, the actual number of plates is unknown. as no catalogs exist, the real number of plates is very difficult to estimate, but for sure the places that we visited have more than 2 million photographic plates and/or negatives in their collections. we plan to continue these efforts, now with emphasis on german astronomical plate collections. figure 5: example of digitized film from sao baker super schmidt camera we plan to continue data mining and plate digitization in astronomical photographic plate/negative archives is as follows. (i) digitize the plate archives using a fast and transportable scanning device, as described above. this scanning method is fast and inexpensive. these are important considerations, as the archives are scattered and there are very large numbers of plates. (ii) create electronic catalogs. (iii) include these catalogs into search programs like wfpd, operated by our bulgarian colleagues (e.g. tsvetkov et al., 2005, and tsvetkov, 2009). acknowledgement the scientific part (investigation of long–term behaviour of high–energy sources) of the study is linked to the grant 13-33324s provided by the grant agency of the czech republic (gacr). the analyzes of astronomical plates are more recently supported by the gacr grant 13-39464j. references [1] hudec, r. et al., acta polytechnica, ibws2011 proceedings, 1(52), 2012. [2] hudec, r. astrophysics with astronomical plate archives, in exploring the cosmic frontier: astrophysical instruments for the 21st century. eso astrophysics symposia, european southern observatory series. edited by andrei p. lobanov, j. anton zensus, catherine cesarsky and phillip j. diamond. series editor: bruno leibundgut, eso. isbn 978-3-540-39755-7. published by springerverlag, berlin and heidelberg, germany, 2007, p.79 [3] hudec, r., an introduction to the world’s large plate archives, acta historica astronomiae, vol. 6, p. 28-40, 1999. [4] robbins, l.; osborn, w., preserving astronomy’s photographic legacy: current state and the future of north american astronomical plates. asp conference series, vol. 410, appendix b and c. edited by wayne osborn and lee robbins. san francisco: astronomical society of the pacific, 2009. 319 introduction the plate archives transportable digitizing device general picture two recently found large us archives with possible major impact on hea summary acta polytechnica ctu proceedings doi:10.14311/app.2016.3.0060 acta polytechnica ctu proceedings 3:60–64, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app numerical method for estimation of tensile load in tie-rods kseniia riabova∗, luca collini, rinaldo garziera university of parma, parco area delle scienze, 181/a, parma, italy ∗ corresponding author: kseniia.riabova@studenti.unipr.it abstract. the work introduces a two-phase method for determination of axial loads in tie-rods. the method described here consists of an experimental activity and an automated numerical calculation. the influence of considering an elastic winkler-type bed to model the tie-rod constraint inside the wall has been investigated. the algorithm used for calculation involves a solution of a functional minimization problem, where the tensile load and the stiffness of elastic foundation at the edges are used as optimization parameters and the error function, which describes the deviation between the frequencies measured and those calculated using finite element method, is minimized. qualitative analysis of the results showed a significant reduction of the error compared to models with different boundary conditions. the method showed to be conservative for the strength evaluation of the rods, because the optimal values of tensile loads appeared to be higher than the load in perfect encastre conditions. keywords: tie-rods, non-destructive technique, numerical method. 1. introduction tie-rods are structural elements used in a wide range of civil constructions. one of purposes they serve is to provide support for masonry arches and vaults in ancient buildings (e.g. churches, castles, etc.), which are known to lurch and founder course of time. tierods are subjected to axial tension and, thus, help the building resist occurring lateral loads exerted by structural elements, walls and facades. over the years deformation of masonry walls and some displacements in the building may cause significant changes in the axial loads of tie-rods. in the extremes this can lead to either failures in tie-rods structural integrity, or to laziness of tie-rods that stop carrying out their duty when they loose loads. that is why regular monitoring of tie-rods condition is of a great importance. the goal of the presented work is to determine axial loads acting on the tie-rods of the fifteenth century casa romei located in ferrara, emilia-romagna, italy, and thus to evaluate their functionality and reliability. an approach based on an experimental investigation and a numerical method is introduced to serve this purpose. 2. previous studies some attempts have been made throughout the years to develop an appropriate non-destructive technique for and indirect estimation of the tensile load in tierods. previous studies [1–3] report a method that combined static and dynamic force identification. tierods were modeled as simply supported euler beams with rotational springs of similar stiffness added on each edge. the stiffness of the spring and the force were the two unknowns obtained from the system of equations, built with a static equation for deflection and a dynamic equation for natural frequencies. another study [4] introduced a static approach for force identification. the experiment consisted of measuring three vertical displacements and strains variations at three sections of the tie-rod under a concentrated load; in [5] an algorithm to identify the axial tensile force in ancient tie-rods by using the first three natural frequencies is developed. the tie-rod was modelled as an euler beam of uniform cross-section, neglecting the shear deformation and rotary inertia, and was assumed to be simply supported at the ends with additional rotational springs. recently, maes et al. [6] introduced a method that enables definition of axial loads in slender beams with unknown boundary conditions, taking into account affects of rotational inertia of the beam and masses of sensors. however, it requires data from five or more sensors along the length of the beam to determine all the introduced unknown of the inverted problem. a similar technique of the axial force identification was developed by li et al. [7] focusing on studies of euler-bernoulli beams and takes into account bending stiffness effects. rebecchi et al. [8] established an analytical method of processing experimental data from five instrumented sections of a prismatic slender beam, which showed excellent results in estimation of the axial load in tie-rods. the method does not require any exact value of effective length of the beam, but neglects both rotary inertia and shear deformations effects in the solution for beam vibrations. for cases of similar beams, tullini et al. [9–11] proposed a static method of axial force identification. the analytical algorithm makes use of any set of experimental data 60 http://dx.doi.org/10.14311/app.2016.3.0060 http://ojs.cvut.cz/ojs/index.php/app vol. 3/2016 estimation of tensile load in tie-rods represented by flexural displacements or curvatures measured at five cross-sections of the beam subjected to an additional concentrated lateral load. again, gentilini et al. [12] developed a procedure that combines dynamical testing with fem simulations using added masses. the method was tested out for tie-rods of various length and load intensity, showing reliable results. livingston et al. [13] identified the tensile force in prismatic beams of uniform section by using modal data and assuming rotational and vertical springs at each end of the beam. shear deformation and rotary inertia were neglected (according to the euler beam model). 3. estimation method in the present study authors intend to introduce a method of health monitoring of tie-rods inserted in ancient monumental masonry buildings. the method consists of an in-situ experimental activity and an automated two-parameter optimization algorithm which allows to evaluate the axial load in tie-rods with a good precision. the technique is basically a simply executed frequency-based identification method that allows to minimize the estimation error. about six natural frequencies can be easily identified with a simple test by experimentally measuring the frequency response functions (frfs) of the tierods with instrumented hammer excitation. further a parametric finite element model of a tie-rod is developed in the fem analysis software abaqus. the part of a tie-rod hidden inside the masonry wall is assumed to have a length of 0.2 m and the constraints given to this part are modelled as elastic foundations (winkler bed). the tensile force and the stiffness of the foundation are chosen as the unknown parameters. in some cases the length of the rod inserted inside the masonry wall can be also assumed as unknown. the sought axial load, as it is explained subsequently, is obtained through an optimization routine that minimizes the deviation between corresponding experimentally determined eigenfrequencies and those calculated by fem. the authors have previously tested the present technique with some variations over different buildings and situations [14–16]. the novelty introduced in the present work is the optimization with respect to the stiffness of the winkler bed type boundary conditions. 3.1. experimental part the tie-rods investigated during the case study have been installed at different points along the lifespan of casa romei (see fig. 1), therefore they differ in dimensions, cross-section shapes and material condition. measurements of geometrical characteristics and of natural frequencies were performed for each reachable tie-rod of the building. the natural frequencies were obtained from the analysis of response to dynamical excitation applied to tie-rods in horizontal plane for a more accurate estimation of the axial load. since figure 1. inner yard of casa romei. figure 2. a typical frf plot (acceleration amplitude vs. frequency). the studied rods are much stiffer in the vertical plane, tests in vertical direction were not performed. the used instrumentation was composed by tools listed in tab. 1. the signal from accelerometers was handled by the dynamic signal acquisition module national instrument (ni) 9234 and acquisition software developed in labview and further elaborated in matlab resulting in the frequency response functions (frf) for each tierod. frf for a ground floor tie-rod pt4 is presented in fig. 2. another important part of the experimental activity was to measure accurately the geometric characteristics of tie-rods, i.e. cross-section dimensions and lengths. for the further study it was assumed that the cross-section remains the same along all the length of a rod. the tested rods are crafted in iron for which the characteristics vary insignificantly, thus the material data was assumed as follows: e = 2 × 1011 n/m2, ρ = 7850 kg/m3, υ = 0.3. the main parameters e and υ appear in the frequencies under the square root, so even some variation in these values can only have a minor influence on the natural frequencies. for the further analysis first six natural frequencies were identified for each tie-rod by means of peakpicking from frfs. six eigenmodes are providing 61 k. riabova, l. collini, r. garziera acta polytechnica ctu proceedings tool manufacturer model sensitivity impact hammer brel&kjaer 8202 4pc/n accelerometer pcb 356a01 0,01 v/(m/s2) accelerometer pcb 356a01 0,01 v/(m/s2) table 1. list of experimental tools. an over constrained problem, thus the higher modes were not taken under consideration in this research. besides that, identification of higher modes might appear inaccurate due to larger possible measurement errors. 3.2. numerical part since the extremities of tie-rods are built into the masonry walls in a way that hardly leaves a possibility to be certain about the end constraints, as a first iteration, boundary conditions for the edges were modeled as plain fixed and/or simply supported. in both cases for any range of optimization parameters (length and force) and weight coefficients the error function reaches minimum every time for the minimal length in the range, i.e. for the measured length. this means that the minimum is forced. investigation of the function in symbolic mathematics package maple 13.0 showed good correspondence of the results obtained from fem with analytical models of a string and a beam under axial load. also the error function has an absolute minimum that lies however below the length measured. this kind of behavior proves that the bcs are irrelevant. the real conditions on the edges of tie-rods are unknown and may vary from one rod to another, because they are fixed inside masonry walls in a way that is hard to determine using non-destructive techniques. however, certain attempts can be made in order to develop a generalized parametric model suitable for numerical determination of loads in tie-rods. in order to determine the loads, tie-rods are modeled in fem software abaqus using 3d beam elements that incorporate timoshenko beam theory (b31), which allows to take into account shear deformation and rotational inertia effects. the cross section is assumed to be constant along each rod. as a first step a pretension load n is applied to the beam and as a second step modal analysis is performed. since the value of the pretension load is the subject of search, it has to be tuned to make the results of modal analysis in fem reach a good agreement with experimentally defined frequencies. however, manual tuning of the load is not considered to be an option in this case, because there is no certainty about the boundary conditions at the beam edges. this problem is overcome by means of a parametric fem model and an optimization tool for computing. in fem model the measured length of the tie-rod is represented by 50 beam elements; boundary conditions are assumed to be an elastic bed and, thus, are represented by 5 spring elements on each edge of the beam, acting in the direction of considered vibrations. also the displacement in direction x (tie-rod axis) is restrained. the sketch of the assumed model is displayed in fig. 3. stiffness of each spring element together with the sought load are chosen as optimization parameters in the problem stated. the optimization program requires a set of experimentally obtained frequencies, ranges of stiffness and force with sizes of steps for each and also a set of weight coefficients for frequencies, that define importance of the corresponding frequency for the analysis. combinations of parameters form a grid and the program launches abaqus input file for each nod of this grid and then extracts and filters natural frequencies. vibrations in z direction only are of our particular interest due to testing that was carried out in the horizontal plane. optimization criteria for the problem is represented by the residual error between experimentally defined frequencies and those obtained from fem analysis. the error e is calculated according to the following formula: e = √√√√ n∑ k=1 p2k(f exp k − f f em k )2 (1) where pk is a weight arbitrary given to each frequency. the optimal values of stiffness of the elastic bed and of axial force deliver minimum value of the error function. the optimization process starts from a reasonable value of the load which can be predicted as an optimal load for a simplified model with fixed ends, i.e. for infinite stiffness of the elastic bed at the constraints. the second step to be done is now to vary the stiffness of the spring bed, and optimize their value to minimize again the error. this implies a refinement process of the parameters along what we call zooming technique [15]. operatively, the algorithm defines a matrix of values and refines grid of parameters to be solved just where local minima of the error as defined in eq. 1 are found. a typical graphic of the surfaces representing the error distribution is presented in eq.4. here the error values are plotted with respect to stiffness and axial load, for a specific tie-rod. the free length l, and length of lf (in fig. 3) are kept constant. the clear minima of the error functional can be observed. 62 vol. 3/2016 estimation of tensile load in tie-rods figure 3. sketch of the computational model. figure 4. error as function of bed stiffness and axial load for different sets of weight coefficients. the error which represents the difference between numerical and experimental frequencies is calculated from the eq. 1. each of the differences (in hz) between couples of frequencies is multiplied by a weight pk in order to allow us to assign higher (or lower) importance to some frequencies rather than the others. the set of weights is then arbitrary chosen, but generally, the first frequencies, or just the fundamental one, have greater importance. a sensitivity analysis showed that the minima of the error functional eq. 1 with respect to the sought tensile load is not influenced by the weights set (see fig. 4). the variation of the error is depicted in fig. 5, wherefrom it can be observed that the minimum value of the error is reached for the optimal elastic bed stiffness and is reached for the up to higher load compared to the encastre boundary conditions, thus, the method reveals to be conservative for the strength evaluation of the rods. the values of error vary from one tie-rod to another, this shows that the model with winkler bed boundary conditions fits better for some of the rods, however these errors were less than those calculated for ordinary fixed or simply supported boundary conditions. for a couple of tie-rods though a perfect encastre boundary condition at the walls appeared to be the best approximation which was demonstrated by an improvement in the error value compared to the case when a set of springs was modelled at their extremities. figure 5. influence of the elastic bed stiffness on the minimal error. 4. conclusions in this work an optimization process is shown and discussed in determining the axial load in structural iron tie-rods. the influence of considering an elastic winkler-type bed to model the tie-rod constraint into the wall has been investigated. this constitutes the major novelty of this work, with respect to the previous studies. structural assessment of the overall building stability is enabled by the described method throughout the identification of axial loads in tie-rods via experimental and numerical activity. in particular, the introduced computational model is able to take into account even unknown boundary conditions that are typical for tie-rods installed in ancient buildings. furthermore, from the simulations it was concluded that the variation of tensile load shifts the set of frequencies (higher the load, higher the frequencies), while the change in elastic bed stiffness changes the distance between natural frequencies. and finally, the conclusion about the method convergence can also be made based on the fact that different sets of weight coefficients bring to minor changes in the sought axial load in tie-rods. references [1] c. blasi. sulla determinazione del “tiro” nelle catene mediante prove statiche e dinamiche. atti iii congresso nazionale assirco 1988. 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[8] g. rebecchi. estimate of the axial force in slender beams with unknown boundary conditions using one flexural mode shape. journal of sound and vibration 332:4122–4135, 2013. doi:10.1016/j.jsv.2013.03.018. [9] n. tullini. dynamic identification of beam axial loads using one flexural mode shape. journal of sound and vibration 318:131–147, 2008. [10] n. tullini. bending tests to estimate the axial force in tie-rods. mechanics research communications 44:57–64, 2012. [11] n. tullini. bending tests to estimate the axial force in slender beams with unknown boundary conditions. mechanics research communications 53:15–23, 2013. [12] c. gentilini. nondestructive characterization of tie-rods by means of dynamic testing, added masses and genetic algorithms. journal of sound and vibration 332:76–101, 2013. doi:10.1016/j.jsv.2012.08.009. [13] t. livingston. estimation of axial load in prismatic members using flexural vibrations. journal of sound and vibration 179:899–908, 1995. doi:10.1016/j.jsv.2012.10.019. [14] m. amabili. estimation of tensile force in tie-rods using a frequency-based identification method. journal of sound and vibration 329:2057–2067, 2010. [15] m. amabili. a hybrid method for the nondestructive evaluation of the axial load in structural tie-rods. nondestructive testing and evaluation 26(2):197–208, 2011. [16] l. collini. load and effectiveness of the tie-rods of an ancient dome: technical and historical aspects. journal of cultural heritage 16:597–601, 2015. 64 http://dx.doi.org/10.1016/j.ijrmms.2008.06.005 http://dx.doi.org/10.1016/j.jsv.2013.03.018 http://dx.doi.org/10.1016/j.jsv.2012.08.009 http://dx.doi.org/10.1016/j.jsv.2012.10.019 acta polytechnica ctu proceedings 3:60–64, 2016 1 introduction 2 previous studies 3 estimation method 3.1 experimental part 3.2 numerical part 4 conclusions references 226 acta polytechnica ctu proceedings 2(1): 226–229, 2015 226 doi: 10.14311/app.2015.02.0226 when a nova becomes old a. ederoclite1, c. tappert2, l. schmidtobreick3, n. vogt2 1centro de estudios de f́ısica del cósmos de aragón (cefca), plaza san juan 1, planta 2, 44001, teruel, spain 2instituto de f́ısica y astronomı́a, universidad de valparaiso, av. gran bretaña 1111, valparaiso, chile 3european southern observatory, alonso de cordova, santiago, chile corresponding author: aederocl@cefca.es abstract here we present the preliminary results of a project aimed at unveiling the nature of classical novae decades after their eruption. the ultimate goal of this project is to describe the population of cataclysmic variables which give rise to nova explosions. so far, in four years of observations, we have concentrated on novae in the southern hemisphere, where we increased by 100% the amount of objects spectroscopically confirmed and increased by 1/5 the amount of objects with known orbital period. keywords: cataclysmic variables classical novae optical spectroscopy photometry. 1 introduction classical novae (cne) are the result of a thermonuclear runaway on the surface of a white dwarf which is accreting mass from a less evolved companion (for a review, see bode & evans, 2008). it is common to define cne as a “subclass of cataclysmic variables” which is a fair statement in both historical and observational terms but it is not quite physically appropriate: a cnexplosion is a phase during the life time of a cataclysmic variable (cv). according to the “hibernation scenario” (shara et al., 1986), post-nova cvs are expected to go through a phase of high mass transfer (triggered by the irradiation of the hot wd) and, eventually, the system gets to separate and spends 90-99% of the inter-explosion time in a quiescent “detached” state (prialnik 1986). it is clear that the use of the term “quiescence” for an oldnova system is very vague and it is the reason why it is used in quotes throughout this work. for a review on old novae, see pagnotta (2013). this project started four years ago trying to answer some fundamental questions: • how does the population of old novae look like? • how does the population of old novae compare to cvs? • do short-period old novae exist in a significant number? • what is the role of magnetic fields on the cnoutburst? despite their brightness at maximum light, the study of novae in “quiescence” is still largely incomplete (again, see pagnotta, 2013). it is important to note that, before this project started, only 39 old novae (out of the 204 novae which exploded before 1980) had a measured orbital period. so far, our project has focused mostly on the southern hemisphere (i.e. targets with δ < 20◦), where, before the start of our project, out of 153 reported novae, 34 lacked identification of the target in “quiescence”, 9 were confirmed but their period was unknown and 24 had a known orbital period. the remaining 86 did not even have a candidate. here we describe the methodology and some early results related with the population which has been recovered so far. for the description of selected targets, see schmidtobreick & tappert (2013) and tappert et al. (2013c) . 2 the methodology our project consists of three phases: the photometric identification of the nova candidate, the spectroscopic confirmation and the period determination through time-series spectroscopy. the target list has been derived mainly from the downes et al. (2005) catalogue. we selected those novae which had exploded before 1980. the selection is made to allow some time after the nova explosion for the characteristics of the underlying cv to become accessible. 226 http://dx.doi.org/10.14311/app.2015.02.0226 when a nova becomes old 1 2 3 4 5 10 20 50 porb(h) 0 2 4 6 8 10 12 14 16 18 this project, n=8 periods of novae before 1980, n=39 1 2 3 4 5 10 20 50 porb(h) 0 5 10 15 20 25 all periods of novae n=78 figure 1: left: period histogram of novae which exploded before 1980. the black histogram is the addition from this project. right: period histogram of all novae (i.e. including those which exploded after 1980). the dotted vertical lines refer to the period gap and the dashed vertical line to the period minimum. we do not include in our sample recurrent novae (rne). these represent thermonuclear explosions analogous to cne but which occur on much shorter recurrence times. the large majority of those rare stars (only a dozen is known in our galaxy) harbour well-evolved secondary stars which sets them somewhat apart from other cvs. as our project primarily originated with the idea of testing the predictions of the hibernation scenario, rne are not included in our analysis. we obtain ubvr photometry of the region of sky where a nova has been reported. here we are taking advantage of the fact that an old nova, as any cv, has a spectrum which is the sum of three components (the white dwarf, the main sequence companion and the accretion disk). this is the same discovery strategy which is being used in large photometric surveys, like sdss (see szkody, 2013). we obtain spectroscopy of those targets whose location in the colour-colour diagram, does not follow the one of the stars. this is “weighted” by the distance from the reported position (i.e. not necessarily the bluest object in the field is the first candidate for follow up spectroscopy, in case there is another object which is closer to the reported position). low-resolution spectroscopy (r = λ ∆λ ' 500) is enough to resolve cvs emission lines and to cover a spectral range (∼4000–9000å) which includes the most relevant lines for cv classification. time-series is obtained once the target has been spectroscopically confirmed. in some cases, time-series photometry is carried out (e.g. tappert et al., 2013b) but, in general, we perform time-series spectroscopy with r = λ ∆λ ' 2500 in the hα region. in order to keep a log of the observations and share the results within a group which is geographically spread over two continents, an internal webpage is being maintained. here we keep a list of our targets, observing logs, and description of the results for each target. 3 results 3.1 the period distribution the left panel of fig.1 shows the histogram of the pre1980 novae. in grey, the periods that can be obtained in the literature and, in black, the addition due to this project. the right panel of fig.1 shows the histogram of all the novae (also the ones which exploded after 1980). the periods of post-1980 novae are from ritter & kolb (2003). qualitatively, the two histograms are very similar: there are few sources with periods within the period gap and a handful of sources below the period gap. there is a “clustering” of sources right above the period gap (just in the period region where sw sex stars belong, see schmidtobreick & tappert, 2013) and less sources at longer periods. this is an interesting result by itself, since our sample is selected on the basis that the targets are back in “quiescence”. the only significant difference is the number of objects found in the 5 – 10 hours period range, which can be easily ascribed to small-number statistics. as we already mentioned, very few objects are within (or below) the period gap. this is consistent with old novae being in a high-mass transfer phase, as predicted by the “hibernation” scenario. it is also worth noting that the overall shape is consistent with the modelling by townsley & bildsten (2005), which assumes that novae occur mostly in high mass accretion rate systems. 227 a. ederoclite et al. 3.2 the role of magnetic fields as we mentioned in the previous section, there are very few objects with periods <3 hours, and, as it can be seen from fig.2, these objects are mostly magnetic. among the 47 pre-1980 old novae with a measured period, only v1500 cyg (1975) and cp pup (1942) are catalogued as am her stars and have periods of 3.36 hours and 1.68 hours, respectively. 0 5 10 15 non magnetic systems am her 1 10 porb(h) 0 5 10 15 non magnetic systems ip 0 5 10 15 non magnetic systems dq her figure 2: the histograms of the magnetic systems which are observed to have harboured a nova explosion. in white, for comparison, the histogram of nonmagnetic systems. in the upper panel, the am her, in the middle panel the dq her systems and, in the lower panel the intermediate polars. the classification is either from ritter & kolb (2003) or from tappert et al. (2012,2013a). although the number of magnetic systems is still small (hence we cannot draw firm conclusions) it is promising that polars and intermediate polars are located in a similar region of the period histogram as in the “classical” cv period histogram. moreover, as mentioned by tappert et al. (2013b), 7 out of 10 novae in the gap may be magnetic. this is in agreement with the population model by townsley & bildstein (2005), who suggest that novae within the period gap are mostly magnetic. araujo-betancor et al. (2005) derive that the total ratio of known magnetic to nonmagnetic cvs is ∼22% (at all periods). the high number of magnetic cvs within the gap, if confirmed, may imply a different evolution through the gap between magnetic and nonmagnetic systems. 3.3 the explosion amplitude from downes et al. (2005), one can derive that the average explosion amplitude is 12.57±2.47 mag. combining our data with data from downes et al. (2005), ritter & kolb (2003), duerbeck (1987), one can try to compare the explosion amplitude (i.e. the difference between the magnitude at maximum and the magnitude in “quiescence”) with the rate of decline (see fig.3). this has already been done by vogt (1990). in fig.3, we show the data from vogt (1990) plus data from our project and post-1980 novae from ritter & kolb (2003). it is interesting to note that our data points (the red squares) fall in the same region as the others. the straight line is the fit from vogt (1990) to the points in his paper and is still valid with the inclusion of the new points. −0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 log t3 4 6 8 10 12 14 16 18 20 d e lt a m figure 3: rate of decline (as measured by the t3 parameter) vs. explosion amplitude (difference between the magnitude at minimum and the magnitude at maximum). the red squares are from this project. the line is not a fit to the present data-set but to the data presented in vogt (1990). the two objects which seem to deviate from the relation are v1500 cyg (1975) and v458 vul (2007). the first one, as mentioned in the previous section, is an am her star. the second one showed a very irregular light curve, with three repeated peaks. these two examples suggest that the scatter in this relation is likely due to two independent causes: on one hand, the magnitude at “quiescence” depends on the structure of the accretion disk (if any) and the magnitude at maximum light (and the rate of decline) may be misestimated due to a series of factors (e.g. bad sampling or dust formation) which are quite typical in novae. warner (1987) has already 228 when a nova becomes old suggested that, for a better comparison of the explosion amplitude, one should take into account, at least, the orbit inclination of the system. another important caveat is that some old novae are observed to decrease their “quiescent” brightness (see e.g. vogt 1990, and references therein, or johnson et al. 2013) and, therefore, fig.3 should be considered in a more dynamical context. 4 summary and future perspectives we have presented a project aimed at the recovery, classification and characterisation of classical nova systems decades after the explosion. during this project, which started four years ago, we have already significantly increased the number of identified and spectroscopically confirmed classical novae (novae without an identified progenitor decreased by 24% and the number of spectroscopically confirmed increased by 100%). finally, we have also increased by 1/5 the number of pre-1980 oldnovae with known orbital period. we plan to finalise the southern part of our project during the next two years and to focus on targets in the northern hemisphere. once the program will be completed, this sample will be crucial in the study of the evolution of classical novae towards quiescence, the study of cn-progenitors as a subclass of the cnpopulation and the development of of the “nova populations” framework (della valle 2002 and references therein). for better sharing the results of our project, we are planning on providing access through the virtual observatory to the reduced data, thus sharing a database with basic parameters (position, brightness, type,. . . ), the spectra used for characterisation and finding charts. this data release will represent a significant legacy for the cv community. acknowledgement the cefca is funded by the fondo de inversiones de teruel, supported by both the government of spain (50%) and the regional government of aragón (50%). this work has been partially funded by the spanish ministerio de ciencia e innovación through the pnaya, under grants aya2006-14056 and through the icts 2009-14. ct and nv acknowledge financial support by fondecyt regular grant 1120338. references [1] araujo-betancor, s. et al.: 2005, apj, 622, 589 [2] bode, m. & evans, n.: 2008, “classical novae”, 2nd edition, cambridge astrophysics series, no 43, cambridge: cambridge university press, 2008 [3] della valle: 2002, aipc, 637, 443 [4] downes, r. et al.: 2005, jad, 11, 2 [5] duerbeck, h.: 1987, space sci. rev. 45, 1 doi:10.1007/bf00187826 [6] johnson, c. et al.: 2013, apj, 780l, 25 [7] pagnotta, a.: 2013, this workshop [8] prialnik, : 1986, apj, 310, 222 [9] ritter h., kolb u. 2003, a&a, 404, 301 doi:10.1086/164762 doi:10.1111/j.1365-2966.2012.21054.x [10] schmidtobreick, l. & tappert, c.: this workshop doi:10.1093/mnras/stt139 [11] shara, m. et al. 1986 apj, 311, 163 doi:10.1093/mnras/stt1747 [12] szkody, p.: 2013, this workshop [13] tappert, c. et al.: 2012, mnras, 423, 2476 doi:10.1086/430594 [14] tappert, c. et al.: 2013a, mnras, 431, 92 doi:10.1086/168866 [15] tappert, c. et al.: 2013b, mnras, 436, 2412 [16] tappert, c. et al.: 2013c, this workshop [17] townsley, d. & bildsten, l.: 2005, apj, 628, 395 [18] vogt, n.: 1990, apj, 356, 609 [19] warner, b.: 1987, mnras, 227, 23 229 http://dx.doi.org/10.1007/bf00187826 http://dx.doi.org/10.1086/164762 http://dx.doi.org/10.1111/j.1365-2966.2012.21054.x http://dx.doi.org/10.1093/mnras/stt139 http://dx.doi.org/10.1093/mnras/stt1747 http://dx.doi.org/10.1086/430594 http://dx.doi.org/10.1086/168866 introduction the methodology results the period distribution the role of magnetic fields the explosion amplitude summary and future perspectives acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0073 acta polytechnica ctu proceedings 4:73–79, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app cfd simulation of upward subcooled boiling flow of freon r12 tomas romsy∗, pavel zacha faculty of mechanical engineering, czech technical university in prague, technická 4, prague 6, 166 07, czech republic ∗ corresponding author: tomas.romsy@fs.cvut.com abstract. subcooled flow boiling under forced convection occurs in many industrial applications to maximize heat removal from the heat source by the very high heat transfer coefficient. this work deals with cfd simulations of the subcooled flow boiling of refrigerant r12 solved by code ansys fluent r16. the main objective of this paper is verification of used numerical settings on relevant experiments performed at debora test facility. also comparisons with previously provided simulation on nri rez are presented. data outputs from this work are basis to subsequent calculations of steam-water mixture cooling of pb-li eutectic. keywords: cfd simulation, subcooled boiling, freon r12. 1. introduction subcooled flow boiling can solve a wide range of high power thermal challenges in the following areas: central processing unit (cpu) and computational application specific integrated circuits (asics) [1, 2], ultra-high brightness light emitting diodes (leds) and lasers sources [3], automotive power electronics [4], or avionics [5], where research findings have made possible the development of a new generation of cooling hardware, which promises order of magnitude increase in heat dissipation compared to today’s cutting edge cooling schemes. subcooled flow boiling is expected to realize the high heat flux cooling for electronic devices [6, 7], or for high heat flux cooling in microgravity [8]. subcooled flow boiling also plays a significant role in case of cooling of plasma facing component of thermonuclear reactors. the subcooled flow boiling is a high complex form of two-phase liquid flow in which the single-phase flow merges into the two-phase flow and back to the singlephase flow. a local appearance and disappearance of the two-phase flow allow the use of the latent heat of vaporization for a heat removal from the surface into the liquid without causing the boiling crisis. in comparison with a single-phase fluid it represents the efficient way to improve heat transfer from the wall to the flowing liquid. therefore, it provides an ability of the sufficient heat transfer at high heat fluxes. detailed investigation of this phenomenon can help for its better understanding to make wider improvement. one of the perspective application of subcooled flow boiling is so-called cold trap facility. this device serves for corrosion impurity separation from the metal eutectic (pb, pb-li), whereas the separation is carried out by the cooling of the eutectic down to the required temperature. this facility considered in centre of advanced nuclear technology project (canut) [10] provides sufficient heat removal from eutectic across the wall to the subcooled steam-water mixture (cooling loop) and subsequently to the tertiary water loop providing ultimate heat removal. part of the cold trap design is cooling concept which strongly depends on steam-water mixture behaviour. description of cooling loop behaviour is provided with the aid of computational fluid dynamics (cfd) code ansys fluent. verification of suitable numerical settings and computational methodology should be determined in the first step. selected experiments on the debora facility were used for this purpose. a detailed description of this device can be found in [11] and [9]. debora project at cea grenoble (french alternative energies and atomic energy commission) was focused on determination of subcooled boiling character by using refrigerant dichlorodifluoromethane (r12). provided experiment outputs contain especially void fraction, gas velocity, temperature of the fluid and the bubble size. because both freon r12 and water have relative physical similarity, we can use this freon for his lower operating parameters in comparison with water. this leads to a much simpler and safer measurements. the geometry of the debora facility is shown in figure 1. it is the vertical heated tube with inner diameter of 19.2 mm. freon is heated along the 3.5 m long part of pipe. the rest of 5 m long tube is unheated. it corresponds to 1 m of inlet section and 0.5 m of outlet section. a radial profile of the vapour volume fraction and its speed at the end of the heated part were measured by optical probe. additionally, the bubble size profiles were available for this area. axial wall temperatures were measured using thermocouples, but radial temperatures were not available [9]. cfd model was performed based on the debora facility parameters and the boundary conditions and standard numerical approaches were set according to the experiment conditions. the results from simula73 http://dx.doi.org/10.14311/ap.2016.4.0073 http://ojs.cvut.cz/ojs/index.php/app tomas romsy, pavel zacha acta polytechnica ctu proceedings figure 1. a) sketch of the debora test geometry [9] b) computational model sketch with detail of used grid. measured quantity p [mpa] v [m/s] qw [w/m2] tin [◦c] tsat [◦c] case 1 3.008 0.86 58 260 67.89 94.24 case 2 3.006 0.891 58 260 73.7 94.21 table 1. specified conditions of the chosen experimental cases. tions provided at nuclear research institute rez (nri rez) [12] are included in section 4 for better comparison. main purpose of this comparison are different approaches in calculation which lead to opportunity to compare them. while the previous simulation from nri rez was performed at ansys fluent r13, presented work was processed at ansys fluent r16. earlier release 13 missed necessary subcooling numerical models and these calculations were implemented by individual user defined functions (udf). however, starting from release 15, numerical models describing this phenomenon are implemented internally. the main output of this work is to determine selected parameters (pressure, velocity, temperatures) in the specified geometry position and compare them against the measured experimental data provided from debora test facility and against the previously calculated values from nri rez simulations. 2. model description simulation model is based on geometry of the debora experimental device. the 10° v-cutout, shown in figure 1, was considered due to the symmetry and to reduce the computational time. slightly rough hexahedral grid was tested in the first steps of the calculation. further enhancements were applied after obtaining better knowledge of the calculation behaviour on this grid. also much finer grid was tested. this grid was only the 10° v-cutout of tested geometry for computational time reduction due to a greater number of cells closer to the wall. however, there was a too small angle of the tip in the axis (centre of the tube) and therefore a smooth auxiliary wall had to be applied with tip neglecting. as the result, flow cross-section has been little changed, but hexahedral cells could be used for the whole geometry. the simulation results for this grid were similar to case with 45° v-cutout geometry. the 45° v-cutout grid has been selected for the final evaluation of the results, because no geometry adjustments are used there. also, it is not appropriate to use fine meshes with respect to numerical approach for the used boiling computational model, like accretion of bubbles near the wall etc. this final grid contains 90 000 hexahedral cells with their higher concentration in the wall vicinity, shown in figure 1. cells quality was evaluated for equi-size skew, which does not exceed value of 0.5 and by the aspect-ratio with the maximum of 10.3. two cases with different boundary conditions based on the debora experiments are considered. the first one – case 1 – has a much larger subcooled inlet state then the case 2, but the saturation conditions and heat flux are almost the same for both of them. the main specified conditions of these experiments are given by table 1 and setting of the boundary conditions in the simulation model corresponds with them. usage of these two experimental cases brings the opportunity to try if the simulation model can solve various ratio of boiling states; from subcooled up to almost boiling crisis – departure from nucleate boiling (dnb). the physical values setting for the freon r12 liquid and vapour is based on these conditions. therefore linear functions depending on the temperature were used for density, specific heat capacity, viscosity and thermal conductivity. these quantities were obtained from nist database [13]. 74 vol. 4/2016 cfd simulation of upward subcooled boiling flow of freon r12 setting of the model boundary conditions: • inlet: velocity inlet – the same speed for both phases turbulence intensity – estimated at 3 % • outlet: pressure outlet – 0 backflow volume fraction • heated wall: heat flux the measuring line was created on one symmetry wall directly at the interface between the heated and outlet non-heated section. this line contains 11 radially distributed evaluation points which were created to monitor calculated data. they are situated in the centre of each appropriate cell and their distribution is shown in the figure 2. radial distance in [m] . ........ . . figure 2. distribution of the evaluation points. 3. solver setting the solver setting is mainly supported by the documentation [12, 14–17]. time dependent solution has to be considered with time step set to 0.01 s for each calculated case. the realizable k−� turbulence model with the non-equilibrium wall function has been chosen. this turbulence model is solved for each phase, but the wall function is calculated only for singlephase. the multiphase coupled model was used for the sequential solution of velocity and pressure fields and the least square cell-based method was set for gradients. all discretization schemes were set to 1st order, otherwise the simulation is very susceptible to the formation and disintegration of bubbles along the simulation domain. calculations were carried out on the sixteen intel xeon e5-2660 processors. the complete final calculation took about 8 hours of computer time for each case. setting of boiling in ansys fluent 16: multiphase eulerian – rpi boiling model this boiling model is appropriate to use for subcooled flow, where condition tsat − tbulk > 3 k is fulfilled. model does not calculate the vapour temperature, but it is fixed at the saturation temperature instead. as an alternative, non-equilibrium boiling model can be used with criterion tsat − tbulk ≤ 3 k, where the vapour temperature is included in the solution process. this condition could occur at the higher model levels, where the liquid can be already sufficiently heated close to saturation conditions. however, the rpi boiling model is considered for the next steps of this work. figure 3 shows scheme of the selected submodels for rpi boiling model. figure 3. scheme of the selected sub-models for rpi boiling model. the wall lubrication model (effect of virtual mass) was neglected, because of big impact to the solution grid fineness, which is not very desirable for the model. 4. calculation results the results of performed calculations are compared with data calculated by nri rez (ansys fluent r13) and measured values from the debora experiment given by [12]. the radial profiles of important monitored parameters for both calculated cases are described below. these profiles are obtained from measuring line, for contained evaluation points respectively. dispositional (orthographic) views are also given for contours of selected parameters in the computational area, see figure 12. if they had not been applied, the side view would have showed just the long and thin tube with nothing obvious to see in the 75 tomas romsy, pavel zacha acta polytechnica ctu proceedings computational domain. given the similarity of the boiling behaviour of both cases, only results for case 1 are listed. 4.1. radial temperature profiles unfortunately due to unavailable experimental data for radial temperature profiles, only the comparison with the results provided by nri rez is mentioned. for both calculated cases the evaluated radial profile has a similar course, as shown in figure 4 for case 1 and in figure 5 for case 2. . . . . . . . . . . . . . . . . . . . . . . figure 4. radial temperature profile, case 1. . . . . . . . . . . . . . . figure 5. radial temperature profile, case 2. only slight differences between the temperature values obtained by nri rez and calculation up to 0.78 ◦c in the fluid midstream and up to 1.36 ◦c near the wall for case 1 are seen. as regards the case 2, here is obvious difference between temperature profiles in the midstream. the nri rez results for case 2 are very close to saturation temperature almost across the whole radial profile, except the wall. the character of ctu calculation temperature profile has a decreasing character up to the axis – tube centre (little bit subcooled state). but near the axis the temperature difference does not exceed 0.58 ◦c. 4.2. radial void fraction profile radial profiles for void fraction with measured data and also nri rez results have been compared already. this is shown in figure 6 (case 1) and figure 7 (case 2). in case 1 is seen that all the data are in good agreement near the wall, but the curves have slightly different courses closer to axis – tube centre. in case of this work, such disagreement is probably caused by collecting of non-condensed secondary phase in the midstream and efforts of liquid go closer toward the wall. this can be seen from void fraction rendered along the calculated area in section 4.5, where the results evaluation along the model is described. figure 6. radial void fraction profile, case 1. figure 7. radial void fraction profile, case 2. 76 vol. 4/2016 cfd simulation of upward subcooled boiling flow of freon r12 the greater disagreement with other data is obvious for case 2. the ctu simulation model estimates almost abrupt increase of void fraction near the heated wall pretty well. but in the midstream the void fraction profile is lower than the others. this probably correlates with the results for temperature quite well. for the treatment of such disagreements it would be possible to try a different approach in model setting, or using of non-equilibrium boiling model for a future simulations respectively. 4.3. radial vapour velocity profile vapour velocity radial profile comparison for case 1 is illustrated in figure 8. it can be seen that the ctu and nri rez results have similar character, but compared with the measurement they have a higher drop of vapour velocity in the fluid midstream. this is probably caused by using of non-ideal single phase wall function. this will be the subject of further investigation in a future simulations. figure 8. radial vapour velocity profile, case 1. in case 2, see figure 9, it is well visible that curve of ctu calculation follows the measured data near the wall pretty well. on the other hand, the same problem as in the previous case occurs in the fluid midstream. figure 9. radial vapour velocity profile, case 2. 4.4. radial bubble diameter profile very important parameter in terms of subcooled boiling behaviour is diameter of vapour bubbles. for experiment measured data; the sauter mean diameter is defined as the diameter of a sphere, which has the same ratio of volume and surface area as the base particle (bubble) and the average diameter dg is the diameter of the bubble, which would have the equivalent two-phase flow with the same density of bubbles amount, the same density of interfacial area and the same diameter for all bubbles, see [12]. for case 1, the results are depicted in figure 10. here is evident that the bubble diameters along the radius have similar size and character as the measured data. it means that the bubbles are breaking up near the wall and their diameter is increasing in the centre of the stream. slighter difference in the fluid midstream probably corresponds with the above listed diagram for void fraction. also bubble diameter along the tube for this case is shown in figure 12. figure 10. radial bubble diameter profile, case 1. figure 11. radial bubble diameter profile, case 2. the results for case 2, shown in figure 11, are in very good agreement with experiment – dg and nri rez near the wall. whereas the bubble diameters in the midstream have little bit lower values compared with other data, they are still in very good agreement 77 tomas romsy, pavel zacha acta polytechnica ctu proceedings figure 12. axial liquid temperature, void fraction and bubble diameter, case 1. with experiment – dg, because it is seen, that the character of radial profile have nearly the same course. smaller values for bubbles are probably caused by lower temperature of fluid in the midstream, shown in figure 5. therefore, the question of using the non-equilibrium boiling model arises here. 4.5. axial results profiles figure 12 (case 1) due to similarity of the both cases of boiling behaviour mentioned above is presented for a better understanding, how the fluid behaves along the tube and where boiling occurs. there are orthographic views for the fluid temperature, void fraction and bubble diameters. there is visible, that all of these quantities correlate with each other. for example; the fluid temperature approximately in the middle of the model reaches the saturation condition near the wall and the void fraction begins to occur there. the higher level the fluid stream in tube reaches, the higher temperature in the midstream occurs. it also means, that the higher concentration of void fraction is there. on the profile of bubble diameters, it is shown how to bubbles reach the greater diameters in the higher levels of the tube and how they are trying to get into the centre of fluid stream. considering this view, it should be observed, that for setting of bubbles modelled by yao-morel model, the smallest possible value of the bubbles have to be set. for both cases this value was set to −105 m. 5. conclusion the main goal of this work was to try the possibility to utilize the newly implemented boiling models in cfd solver ansys fluent r16. the debora experiments were chosen for comparison with data received from these simulations. some of the earlier works, such as nri rez (fluent r13), were also tried to simulate boiling experiments by using older version of fluent cfd solver. however, suitable boiling models had to be supplied by user defined functions (udf) in this earlier version. for comparison, how the solver with newly implemented boiling models (fluent r16) stand up against the older version with udf’s, the evaluation of these data compared with ctu is described. two cases (case 1 and case 2) of debora experiments were chosen. only inlet conditions for fluid temperature are different in these two cases. it means that the first one is much more subcooled case then the second one. the comparison between measured data, calculations from nri rez and results of this work are in the high level of agreement. only few small differences occur. the vapour velocity disagreement of both calculations can be primarily caused 78 vol. 4/2016 cfd simulation of upward subcooled boiling flow of freon r12 by numerical approach, or by using of non-ideal single phase wall function respectively. also, for radial void fraction closer to axis the curve have slightly different course compared to experiment and nri rez. this disagreement is probably caused by collecting of non-condensed secondary phase in the midstream and effort of liquid go closer toward the wall. and also in case 2 the void fraction has the lower values in the fluid midstream. therefore the question of using the non-equilibrium boiling model arises here. it will be subject of further research, because it can treat the fact, that for case 2 the midstream fluid temperature could reache the values close to saturation conditions in the measured section. after evaluation of all the results we can say, that the cfd solver with newly implemented boiling models (fluent r16) can be successfully used for problems where the subcooled boiling occurs. but some minor deficiencies are still present. unfortunately due to very complicated and sensitive numerical model used for description of this boiling phenomenon, there are many degrees of freedom at the model settings. also variables that are tracked interact with each other and individual numerical approaches have limited validity. this leads to time-consuming study to solve these minor deficiencies and verifying all of the used hypotheses (and their interaction) by experiments. all the knowledge and experiences described above are currently being used for simulations of more complicated situation of subcooled boiling of water, which occurs for example in pb-li cold trap cooling system, and they are continuously being improved. list of symbols p pressure [mpa] v inlet liquid velocity [m/s] qw heat flux [w/m2] t temperature [°c] tin inlet liquid temperature [°c] tbulk bulk liquid temperature [°c] tsat saturation temperature [°c] % liquid density [kg/m3] cp specific heat capacity [kj/kg k] ν dynamic viscosity [pa s] λ thermal conductivity [w/m k] references [1] s. g. kandlikar. high flux heat removal with microchannels – a roadmap of challenges and opportunities. heat transfer engineering 26(8):5–14, 2005. doi:10.1080/01457630591003655. [2] n. khan, k. c. toh, d. pinjala. boiling heat transfer enhancement using micro-machined porous channels for electronics cooling. heat transfer engineering 29(4):366–374, 2008. doi:10.1080/01457630701825481. [3] a. bar-cohen, p. wang, e. rahim. thermal management of high heat flux nanoelectronic chips. microgravity science and technology 19(3):48–52, 2007. doi:10.1007/bf02915748. [4] f. ramstorfer, h. steiner, g. brenn, et al. subcooled boiling flow heat transfer from plain and enhanced surfaces in automotive applications. asme j heat transfer 130(1), 2008. doi:10.1115/1.2780178. [5] i. mudawar. assessment of high-heat-flux thermal management schemes. ieee transactions on components and packaging technologies 24(2):122–141, 2001. doi:10.1109/6144.926375. [6] j. l. parker, m. s. el-genk. enhanced saturation and subcooled boiling of fc-72 dielectric liquid. international journal of heat and mass transfer 48(18):3736–3752, 2005. doi:10.1016/j.ijheatmasstransfer.2005.03.011. [7] y. madhour, j. olivier, e. costa-patry, et al. flow boiling of r134a in a multi-microchannel heat sink with hotspot heaters for energy-efficient microelectronic cpu cooling applications. ieee transactions on components, packaging and manufacturing technology 1(6):873–883, 2011. doi:10.1109/tcpmt.2011.2123895. [8] k. suzuki, h. kawamura. microgravity experiments on boiling and applications: research activity of advanced high heat flux cooling technology for electronic devices in japan. annals of the new york academy of sciences 1027(1):182–195, 2004. doi:10.1196/annals.1324.017. [9] e. krepper, r. rzehak. cfd for subcooled flow boiling: simulation of debora experiments. nuclear engineering and design 241(9):3851–3866, 2011. doi:10.1016/j.nucengdes.2011.07.003. [10] o. frybort. background report – pb-li17 cold trap. nuclear research centre rez, 2014. [11] j. garnier, e. manon, g. cubizolles. local measurements on flow boiling of refrigerant 12 in a vertical tube. multiphase science and technology 13(1&2), 2001. doi:10.1615/multscientechn.v13.i1-2.10. [12] l. vyskočil. simulation of the subcooled boiling in fluent 13. nuclear research institute rez, 2011. [13] e.w. lemmon, m.o. mclinden, d.g. friend. thermophysical properties of fluid systems. in nist chemistry webbook, nist standard reference database number 69, p. linstrom, w. mallard (eds.). national institute of standards and technology. http://webbook.nist.gov. [14] t. romsy, p. zacha. cfd simulation of subcooled boiling, 2016. erin, the 8th international conference for young researchers and phd students, april 23–25, 2014, blansko, czech republic. [15] ansys, inc. fluent user’s guide, 2013. release 15.0. [16] ansys, inc. fluent theory guide, 2013. release 15.0. [17] ansys, inc. fluent multiphase 15.0 optional lecture 03 – wall boiling models: training materials, 2014. 79 http://dx.doi.org/10.1080/01457630591003655 http://dx.doi.org/10.1080/01457630701825481 http://dx.doi.org/10.1007/bf02915748 http://dx.doi.org/10.1115/1.2780178 http://dx.doi.org/10.1109/6144.926375 http://dx.doi.org/10.1016/j.ijheatmasstransfer.2005.03.011 http://dx.doi.org/10.1109/tcpmt.2011.2123895 http://dx.doi.org/10.1196/annals.1324.017 http://dx.doi.org/10.1016/j.nucengdes.2011.07.003 http://dx.doi.org/10.1615/multscientechn.v13.i1-2.10 http://webbook.nist.gov acta polytechnica ctu proceedings 4:73–79, 2016 1 introduction 2 model description 3 solver setting 4 calculation results 4.1 radial temperature profiles 4.2 radial void fraction profile 4.3 radial vapour velocity profile 4.4 radial bubble diameter profile 4.5 axial results profiles 5 conclusion list of symbols references acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0097 acta polytechnica ctu proceedings 4:97–101, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app quench front propagation in the annular channel jan štěpánek∗, vaclav bláha, vaclav dostál faculty of mechanical engineering, czech technical university in prague, technická 4, prague, czech republic ∗ corresponding author: j.stepanek@fs.cvut.cz abstract. understanding the quench front propagation during bottom core reflooding is crucial for the effective cooling during the loca accident. the results presented in this paper were obtained on an experimental loop with an annular test section. the test section consists of a vertical electrically heated stainless steel tube with outer diameter 9 mm and length of 1.7 m. the heated tube is placed inside a glass tube with the inner diameter 14.5 mm. water mass flux during the reflooding is in the range from 100 kg m−2 s−1 up to 140 kg m−2 s−1 and the initial wall temperature of the stainless steel tube is in the range from 250 ◦c up to 800 ◦c. the presented results show the influence of the initial conditions on the quench front propagation and the complexity of the phenomenon. keywords: quench front, rewetting, loca. 1. introduction cooling of very hot surfaces is still not satisfactorily described area. the main difficulty of this kind of cooling is that coolant can’t reach the hot surface because of a thin vapor layer generated between the hot surface and the coolant. this vapor layer acts as a thermal insulating gap and it prevents effective cooling. heat transfer in this gap is much lower than in rewetted area. as the surface cools down, the vapor layer collapses and coolant rewetts the surface. a high heat transfer rate takes place at this moment. the location where coolant, for example water, rewetts the surface is called the quench front. because of rapid cooling at the quench front, a heat accumulated in front of the quench front is conducted towards the rewetted area. this heat conduction accelerates the cooling of dry surface and it helps the quench front to advance forward [1]. the speed of this rewetting advance is called the quench front velocity. quenching takes place in many technical applications such as steel hardening, cryogenic technology and many others. one of the most crucial applications is a nuclear reactor safety. this phenomenon is dominant for safety during core reflooding phase of loca (loss-of-coolant accident). the temperature of the fuel cladding rapidly rises up (over 1200 ◦c [2]) after blow-down phase. then emergency core cooling system (eccs) starts to pump relatively cold water into the core. as the water reaches the hot fuel rods quenching appears [3]. this complicates effective core cooling. the quenching phenomenon has been solved analytically in many studies based on experimental data [4]. in these studies cooled geometry is usually divided into several regions. the most common approach divides the geometry into two regions. one region is dry region in front of the quench front. in this region a heat transfer coefficient is assumed as zero, i.e. adiabatic boundary condition. the second region is wet or rewetted region, where a heat transfer coefficient is taken as constant number [5]. this problem is then usually solved using wiener-hopf technique or heat balance integral method (hbim). there were many experimental efforts in the second half of 20th century. but many of these experiments were done for short geometries with limited range of initial wall temperatures [6]. the main goal of upcoming experiments is to cover wide range of initial wall temperatures, water mass flow rates and to describe effects of accumulated heat (different wall thickness and uneven tube heating) and geometrical elements such as spacer grids. in this paper a first set of experimental data is presented. 2. experimental loop for the purpose of quench front phenomenon investigation an experimental loop has been built. the loop consists of test section, dc power source, data acquisition system and variable hydraulic circuit. the test section is assembled from inner electrically heated steel tube and from outer glass tube. the steel tube is 1.7 m high and it is equipped with set of k-thermocouples (see tc1, tc2 and tc3 in fig. 1). these thermocouples are in contact with inner surface of the heated tube. the thermocouples in the wall are not calibrated and its precision is then ±2.5 ◦c but the precision of these thermocouples is not important because temperature time gradient is measured. the spacing between thermocouples is 0.5 m. the outer diameter of the steel tube is 9 mm, its wall thickness is 0.5 mm and the tube is made of x6crniti18 stainless steel. the outer glass tube has inner diameter of 14.5 mm and its wall thickness is 1.75 mm. it gives annular flow cross-section of 100 mm2. the test section is situated between lower and upper chamber. in these chambers are situated pressure sensors with accuracy of 0.5 % fso and calibrated class 2 k-thermocouples. initial wall temperature, water mass flow rate, lower chamber 97 http://dx.doi.org/10.14311/ap.2016.4.0097 http://ojs.cvut.cz/ojs/index.php/app j. štěpánek, v. bláha, v. dostál acta polytechnica ctu proceedings pressure, inlet temperature, power input and outer temperature of the glass tube (tcgl in fig. 1) are scanned during the experiment. the water mass flow rate is continuously measured by turbine flow meter. the precision of the flow meter is ±3 %. p tin tc1 tc3 tc2 q + tcgl figure 1. simplified experimental loop scheme. 3. setup the initial wall temperature of the experiment was in range from 250 ◦c up to 800 ◦c with step of 50 ◦c, the initial flow rate was set to 130 kg m−2 s−1 and its actual values are dependent on hydraulic characteristics of the test section during the experiment. this initial flow rate is adjusted in short circuit of the loop. firstly the steel tube is heated up to desired temperature. the thermocouple on the glass tube is used as an indicator of the stationary state. when the temperature of this thermocouple is constant, the reflooding begins. once the water reaches upper chamber, the experiment is finished. collected information are then evaluated on the pc. 4. quench front velocity in experimental studies quench front velocity is usually identified in several ways. one of these approaches is usage of time dependence of measured temperature on the cooled geometry. another approach for example uses optical methods such as high speed cameras. the quench front velocity was obtained as the distance between thermocouples (0.5 m) overcame in measured time. time, when a thermocouple has been rewetted is identified via maximum value of the first derivative of the measured temperature. this derivative value also represents position of the quenching temperature. when all rewetting times for each thermocouple are obtained, the quench front velocity can be calculated. 5. results each experiment was repeated five times. the measured temperatures and other experimental data were interpolated using piecewise cubic hermite interpolation polynominal (pchip). the quenching temperature has been calculated as an average from valid experimental data. in fig. 2 we can see the quenching temperatures for each initial wall temperature level and for each thermocouple. as can be seen this dependence is almost linear. the rewetting temperature of the first thermocouple behind the inlet is highest and for the last thermocouple is this temperature the lowest of all. this temperature drop is caused by two-phase mixture propelled by steam generated at the quench front. this mixture precools dry region of the heated tube and this precooled surface is then rewetted at lower temperature. initial wall temperature also strongly influences the value of the quench front velocity. higher wall temperature generates more steam at the quench front. on the one hand this steam helps to cool down the dry region but on the other hand this steam generates higher pressure drop and this leads to lower quench front velocity. but the most significant slowdown is caused by evaporated fraction of the inlet water mass flux. this dependence can be seen in fig. 3. from fig. 3 we can see, that quench front velocity drop is very significant between initial wall temperature levels 250 ◦c and 300 ◦c. from 300 ◦c this dependence is almost linear. for 300 ◦c the quench front velocity is 7.6 cm s−1 and for 800 ◦c 2.6 cm s−1. this velocity difference represents about 70 % velocity drop in this temperature range. with higher wall temperatures a significant pressure bump can be observed when the water meets the hot surface in the test section. this prompt pressure increase is caused by rapid steam generation at the quench front. the steam moves at high speed along the hot surface and the steam is then slightly overheated. these conditions lead to high pressure drop along the test section and it slows the quench front down. the pressure drop can be observed from initial wall temperature of 400 ◦c and it grows from 2 kpa up to 17 kpa at 800 ◦c wall temperature level. the value of the pressure bump was calculated as a difference between maximum pressure and uninfluenced pressure at lower wall temperatures in the lower chamber. this dependence is plotted in fig. 4. the maximum value of pressure bump at 800 ◦c represents nearly 90 % of pressure drop of the fully reflooded test section. also heat transfer rate at the quench front has been calculated. this heat transfer rate is calculated using first derivative of the wall temperature and material characteristics of the heated tube such as 98 vol. 4/2016 quench front propagation in the annular channel figure 2. rewetting temperature on initial wall temperature. figure 3. quench front velocity on initial wall temperature. heat capacity and electric resistance. using known difference of accumulated heat and known value of electric current a heat transfer rate at the quench front can be enumerated. when the heat transfer rate is known, the heat transfer coefficient can be also obtained from this heat transfer rate and from known quenching temperature and temperature of saturated water. resulting values of heat transfer coefficient are plotted in fig. 5. 6. experiment and real nuclear reactor the heated tube diameter corresponds to dimension of fuel rod of pressurized water reactor and its length is approximately halved. the results from the experiment give us the basis for estimating the quench front velocity as a local process. in a real nuclear reactor this process goes into complex spatial process. on the other hand a material used for the heated tube 99 j. štěpánek, v. bláha, v. dostál acta polytechnica ctu proceedings figure 4. pressure bump on initial wall temperature. figure 5. heat transfer coefficient at the quench front on initial wall temperature. also affects the quench front propagation. the use of zirconium alloy will probably significantly slow down the quench front propagation along the heated surface due to exothermic reaction of the zirconium and steam vapour approximately above 800 ◦c [7]. inclusion of this phenomenon is necessary for experimental efforts at very high temperatures in the future. the heat generation along the tube is constant in this effort but also an uneven heat generation will be investigated in the upcoming experimental work. this unevenness will affect the quench front velocity as well. 7. conclusions the presented results in this paper show the complexity of the quenching phenomenon for one coolant mass flux level and for range of initial wall temperature from 250 ◦c up to 800 ◦c in the annular channel. from these results can be concluded, that quenching velocities and rewetting temperatures are strongly dependent on the initial wall temperature. with increasing initial wall temperature the quench front velocity rapidly decreases. moreover increasing initial 100 vol. 4/2016 quench front propagation in the annular channel wall temperature goes with significant pressure bumps. these pressure bumps strongly influence the hydraulic characteristic of the heated channel and this leads to a big pressure drop in the channel. these effects can lead to unevenness of water flow through the core during loca. the unevenness can lead to cooling water flowing around the central area of the core, where the surface temperature is lower, and the central area of the core can be cooled down distinctly later. on the other hand, geometrical elements, such as spacer grids, create secondary quench fronts moving along the fuel rod [8]. due to these secondary quench fronts, the last rewetted point can’t be clearly determined. in other words, the water presence above the core doesn’t mean that the rest of the core is rewetted and cooled down. all these weaknesses of the quenching phenomenon should be uncovered for better handling of nuclear reactor accident. more experimental efforts will be done in the future. in these efforts a wide range of water mass flux will be covered. references [1] e. elias, g. yadigaroglu. a general one dimensionalmodel for conduction-controlled rewetting of a surface. nuclear engineering and design 42 (1977), 185–194. [2] m. billone, y. yan, t. burtseva, r. daum. cladding embrittlement during postulated loss-of-coolant accidents. u.s. nuclear rerulatory commision nureg/cr-6967, anl-07/04, 2008. http://www.nrc.gov/docs/ml0821/ml082130389.pdf [3] h.c. yeh. analysis of rewetting of a nuclear fuel rod in water reactor emergency core cooling. nuclear engineering and design 34 (1975), 317–322. [4] s.k. sahu, p.k. das, s. bhattacharyya. analytical and semi-analytical models of conduction controlled rewetting: a state of the art review. thermal science 19 (2015), 1479–1496. doi:10.2298/tsci121231125s. [5] s. olek. on the two region model with a step change in the heat transfer coefficient. nuclear engineering and design 108 (1988), 315–322. [6] a.k. saxena, v. venkat raj, v. govardhana rao. experimental studies on rewetting of hot vertical annular channel. nuclear engineering and design 208 (2001), 283–303. doi:10.1016/s0029-5493(01)00356-9. [7] j. belle, m.w. mallett. kinetics of the high temperature oxidation of zirconium. j. electrochem. soc. 101 (1954), 339–342. doi:10.1149/1.2781278. [8] j. stepanek, v. blaha, v. dostal, p. burda. the effect of spacer grid’s elements on the rewetting velocity. in proceedings of international conference on nuclear engineering (icone23). new york: asme, 2015. isbn 978-4-88898-256-6. 101 http://www.nrc.gov/docs/ml0821/ml082130389.pdf http://dx.doi.org/10.2298/tsci121231125s http://dx.doi.org/10.1016/s0029-5493(01)00356-9 http://dx.doi.org/10.1149/1.2781278 acta polytechnica ctu proceedings 4:97–101, 2016 1 introduction 2 experimental loop 3 setup 4 quench front velocity 5 results 6 experiment and real nuclear reactor 7 conclusions references 143 acta polytechnica ctu proceedings 2(1): 143–147, 2015 143 doi: 10.14311/app.2015.02.0143 model atmosphere spectrum fit to the soft x-ray outburst spectrum of ss cyg v. f. suleimanov1,2, c. w. mauche3, r. ya. zhuchkov2, k. werner1 1institute for astronomy and astrophysics, kepler center for astro and particle physics, eberhard karls university, sand 1, 72076 tübingen, germany 2kazan (volga region) federal university, kremlevskaya str. 18, 42008 kazan, russia 3lawrence livermore national laboratory, l-473, 7000 east ave., livermore, ca 94550, usa corresponding author: suleimanov@astro.uni-tuebingen.de abstract the x-ray spectrum of ss cyg in outburst has a very soft component that can be interpreted as the fast-rotating optically thick boundary layer on the white dwarf surface. this component was carefully investigated by mauche (2004) using the chandra letg spectrum of this object in outburst. the spectrum shows broad (≈ 5 å) spectral features that have been interpreted as a large number of absorption lines on a blackbody continuum with a temperature of ≈ 250 kk. because the spectrum resembles the photospheric spectra of super-soft x-ray sources, we tried to fit it with high gravity hot lte stellar model atmospheres with solar chemical composition, specially computed for this purpose. we obtained a reasonably good fit to the 60–125 å spectrum with the following parameters: teff = 190 kk, log g = 6.2, and nh = 8 · 1019 cm−2, although at shorter wavelengths the observed spectrum has a much higher flux. the reasons for this are discussed. the hypothesis of a fast rotating boundary layer is supported by the derived low surface gravity. keywords: cataclysmic variables dwarf novae radiation transfer x-rays individual: ss cyg ≡ bd+42◦ 4189a. 1 introduction ss cyg is one of the brightest cataclysmic variables (cvs), one of the best-studied dwarf nova stars (warner 1995), and was the first cv discovered in x-ray radiation (rappaport et al. 1974). the properties of the x-ray radiation of this close binary system have been extensively investigated, and are observed to be dramatically different in quiescence and in outburst. in quiescence, the x-ray spectrum is hard and can be described by an optically thin hot (kt ≈ 20 kev) plasma with an observed flux ≈ 2 · 10−10 erg s−1 cm−2. in outburst, this hard component decreases by a factor of ten, the plasma temperature is reduced to ∼ 6–8 kev, and an additional soft component appears with a blackbody temperature ≈ 200–300 kk (córdova et al. 1980; mcgowan et al. 2004; ishida et al. 2009). it is commonly accepted that the x-ray radiation of non-magnetic cvs arises in the boundary layer (bl) between the white dwarf (wd) and the accretion disc (pringle & savonije 1979; tylenda 1981; patterson & raymond 1985a, b; kley 1991), which are optically thick at high accretion rates (ṁ > 1016 g s−1) and optically thin at lower accretion rates. the soft x-ray spectrum of ss cyg in outburst was carefully investigated by mauche (2004) using a highresolution spectrum obtained with the chandra letg. he phenomenologically described the observed 40–130 å spectrum by a blackbody with temperature t ≈ 250 kk and numerous broad absorption features of ions of cosmically abundant o, ne, mg, si, s, and fe. the bl luminosity and wd spin were also evaluated in this work. on the other hand, this spectrum looks like the photospheric spectra of super-soft x-ray sources (lanz et al. 2005; rauch et al. 2010; van rossum 2012), so it probably could be described using the spectra of hot stellar model atmospheres. boundary layers possibly rotate with almost keplerian velocities and could have reduced (in comparison with wd) surface gravities close to the local eddington limit. therefore, we consider close to eddington limit models in the present work. here we present our attempt to fit the chandra letg spectrum of ss cyg using such model spectra. we also make estimates of the bl parameters in the context of our model fits. 2 model atmospheres the version of the lte computer code atlas (kurucz 1970; 1993), modified by us to deal with high temperatures (ibragimov et al. 2003; suleimanov & werner 2007), was used to model high temperature at143 http://dx.doi.org/10.14311/app.2015.02.0143 v. f. suleimanov et al. mospheres. in this code, local thermodynamic equilibrium (lte) is assumed and the pressure ionization effects using the occupation probability formalism (hummer & mihalas 1988) as described by hubeny et al. (1994) are taken into account. coherent electron scattering together with the free-free and bound-free transitions of all ions of the 15 most abundant elements using cross-sections from verner & yakovlev (1995) were adopted for the continuum opacity. line blanketing is also included using ∼ 25000 spectral lines from the chianti, version 3.0, atomic database (dere et al. 1997). 140 160 180 200 220 240 260 5.6 5.8 6.0 6.2 6.4 6.6 6.8 g < g edd lo g g t eff , kk 10-5 10-4 10-3 10-2 10-1 1 10 102 0 1 2 3 4 5 6 7 8 9 10 11 12 g r ad / g column density, g cm-2 t eff = 200 kk, log g = log g edd + 0.2 = 6.28 t eff = 200 kk, log g = log g edd + 0.4 = 6.48 t eff = 250 kk, log g = log g edd + 0.2 = 6.67 figure 1: top panel: positions of the computed model atmospheres in the teff –log g plane. the dashed curve demarcates the eddington limit (log g = log gedd). bottom panel: the relative radiation force vs. depth for various model atmospheres. twenty-two model atmospheres with solar chemical composition were computed using the described code. the effective temperatures of the models range between 150 kk and 250 kk with a step of 10 kk. two values of the surface gravity for each effective temperature, namely log g = log gedd + 0.2 and log g = log gedd + 0.4, were used. here log gedd = log(σe σsbt 4 eff/c) = 4.88 + 4 log(teff/10 5 k) is the surface gravity that has an equal radiation pressure force for a given teff , and σe ≈ 0.34 cm2 g−1 is the electron scattering opacity for the assumed solar chemical composition. the positions of the computed models on the teff –log g plane are shown in fig. 1 (top panel). for the considered model atmospheres, a radiation pressure force grad due to spectral lines becomes larger than the surface gravity at the upper atmospheric layers (see fig. 1, bottom panel). to enforce hydrostatic equilibrium, we took a gas pressure equal to 10% of the total pressure (pgas = 0.1ptot) at all atmospheric layers where grad > g. as shown in fig. 2, the computed emergent spectra are dominated by a forest of absorption lines, and have to be convolved with the letg spectral resolution and the interstellar gas transmission to be compared with the observed spectrum of ss cyg (see fig. 3). the spectra of models with different surface gravities are sufficiently different to discriminate them from a comparison with the observed spectrum (see fig. 2, bottom panel). 40 60 80 100 1011 1012 1013 1014 1015 1016 1011 1012 1013 1014 1015 1016 log g = log g edd + 0.2 = 6.28 log g = log g edd + 0.4 = 6.48 t eff = 200 kk h , er g cm -2 s -1 a -1 wavelength, a t eff = 150, 200, 250 kk log g = log g edd + 0.2 h , er g cm -2 s -1 a -1 figure 2: top panel: emergent spectra of three model atmospheres with the same log g = log gedd + 0.2 and various effective temperatures: 150 kk (solid curves), 200 kk (dashed curves), and 250 kk (dotted curves). bottom panel: emergent spectra of two model atmospheres with the same effective temperature (200 kk) and different log g. 3 results the model spectra convolved with the chandra letg spectral resolution ∆λ = 0.05 å were used to fit the observed soft x-ray spectrum of ss cyg. the interstellar absorption (with the hydrogen column number den144 model atmosphere spectrum fit to the soft x-ray outburst spectrum of ss cyg sity nh as a fitting parameter) was also taken into account. the observed spectrum was fitted in the 60–125 å wavelength range because at the shorter wavelengths our model spectra could be incorrect (see next section). the best-fit model parameters are teff = 190 kk, nh = 8·1019 cm−2, and normalization k = 7.82·10−26, and correspond to the models with the lower surface gravity log g = log gedd + 0.2. the reduced χ 2 = 3.9 is relatively large, hence the formal parameter errors are large, too, and we have not attempted to determine them. the best-fit model spectrum together with the observed spectrum is shown in fig. 3. the contours of χ2 on the teff –log nh parameter plane are shown in fig. 4. the normalization can be expressed as k = fr2wd/d 2 where d is the distance to ss cyg and f is the wd fractional area occupied by the bl, which can be expressed as the relative bl extension along the wd surface f ≈ (2πrwd 2hbl)/(4πr2wd) = hbl/rwd. the basic properties of the bl can be derived from the obtained fit parameters. using the same system and wd parameters as used by mauche (2004) — mwd = 1 m�, rwd = 5.5 · 108 cm (therefore, model log g = 8.46), d = 160 pc, and an accretion disk bolometric luminosity in outburst ldisk = 10 35 erg s−1 — we obtain the fractional area of the bl f = 6.3 · 10−2 (5 · 10−3), the bolometric bl luminosity lbl = 1.8 · 1034 (5 · 1033) erg s−1, and the relative bl luminosity lbl/ldisk = 0.18 (0.05), where the best-fit parameter values obtained by mauche (2004) are shown in parentheses for comparison. the spin period of the wd in ss cyg is 12 (9) s as inferred using the relation between the bl and accretion disk luminosities (kluźniak 1987; kley1991) lbl/ldisk = [1 − ωwd/ωk(rwd)]2, (1) where ωk(rwd) is the kepler angular velocity at the wd radius. the accepted model gravity on the wd surface in ss cyg (log gwd = 8.46) is more than two orders of magnitude higher than the obtained bl effective surface gravity log geff = 6.2. the surface gravity of the bl can be reduced only by fast rotation of the accreting matter geff = gwd − ω2blrwd (2) = gwd (1 − [ωbl/ωk(rwd)]2). therefore, a relative bl angular velocity ωbl/ωk(rwd) ≈ 0.98 was obtained using this relation. figure 3: the chandra letg spectrum of ss cyg in outburst (thick black curve) and the best-fit model atmosphere spectrum with teff = 190 kk, log g = 6.2, and log nh = 19.9 (thin red curve). the fitting was performed in the 60–125 å wavelength range. the model spectrum at shorter wavelengths is shown by the dashed red curve. figure 4: position of the best-fit model in the teff –log nh parameter plane and contours of χ 2 = [1.5, 3, 6] χ2min. 4 discussion and conclusion we present here the results of fitting the ss cyg chandra letg spectrum in outburst with model atmosphere spectra. the obtained best fit model spectrum does not describe the observed spectrum at short wavelengths (< 60 å) and in the 82–90 å wavelength region and, therefore, it is not statistically acceptable (χ2/dof = 3.9). these deficiencies can be connected with model shortcomings: the most important ignored effect is atmosphere expansion due to a spectral linedriven stellar wind. this expansion can be significant 145 v. f. suleimanov et al. because grad > g at the outer layers of our model atmospheres (see also van rossum 2012). in addition, non-lte effects could be important (see, e.g., rauch et al. 2010), the chemical composition may differ from solar, and, finally, the atomic data are almost certainly neither complete nor entirely accurate. the second important shortcoming is connected with a likely complicated bl structure with a distribution of effective temperatures and surface gravities over its surface. therefore, a simple one-zone bl model presented here is most probably insufficient and more sophisticated bl models have to be considered. nevertheless, the spectral modeling presented here supports a bl in ss cyg that is, to first approximation, a hot (≈ 190 kk), fast rotating [ωbl ≈ 0.98 ωk(rwd)], narrow (hbl ≈ 0.063 rwd) belt on the wd surface. acknowledgement this work is supported by the dfg sfb / transregio 7 “gravitational wave astronomy” (v.s.) and the russian foundation for basic research (grant 12-02-97006r-povolzhe-a) (r.zh.). c.w.m.’s contribution to this work was performed under the auspices of the u.s. department of energy by lawrence livermore national laboratory under contract de-ac52-07na27344. references [1] córdova, f. a., chester, t. j., tuohy, i. r., & garmire, g. p. 1980, apj, 235, 163 [2] dere, k. p., landi, e., mason, h. e., monsignori fossi, b. c., & young, p. r. 1997, a&ass, 125, 149 [3] hubeny, i., hummer, d. g., & lanz, t. 1994, a&a, 282, 151 [4] hummer, d. g., & mihalas, d. 1988, apj, 331, 794 doi:10.1086/166600 [5] ibragimov, a. a., suleimanov, v. f., vikhlinin, a., & sakhibullin, n. a. 2003, astronomy reports, 47, 186 doi:10.1134/1.1562213 [6] ishida, m., okada, s., hayashi, t., nakamura, r., terada, y., mukai, k., & hamaguchi, k. 2009, pasj, 61, 77 [7] kley, w. 1991, a&a, 247, 95 [8] kluźniak, w. 1987, ph.d. thesis, stanford univ. [9] kurucz, r. l. 1970, sao special report, 309 [10] mauche, c. w. 2004, apj, 610, 422 [11] mcgowan, k. e., priedhorsky, w. c., & trudolyubov, s. p. 2004, apj, 601, 1100 doi:10.1086/380758 [12] patterson, j., & raymond, j. c. 1985a, apj, 292, 535 doi:10.1086/163187 [13] patterson, j., & raymond, j. c. 1985b, apj, 292, 550 doi:10.1086/163188 [14] pringle, j. e., & savonije, g. j. 1979, mnras, 187, 777 doi:10.1093/mnras/187.4.777 [15] rauch, t., orio, m., gonzales-riestra, r., nelson, t., still, m., werner, k., & wilms, j. 2010, apj, 717, 363 doi:10.1088/0004-637x/717/1/363 [16] rappaport, s., cash, w., doxsey, r., mcclintock, j., & moore, g. 1974, apj, 187, l5 doi:10.1086/181378 [17] suleimanov, v., & werner, k. 2007, a&a, 466, 661 [18] tylenda, r. 1981, acta astr., 31, 267 [19] van rossum, d. r. 2012, apj, 756, 43 [20] verner, d. a., & yakovlev, d. g. 1995, a&ass, 109, 125 [21] warner, b. 1995, cambridge astrophysics series, 28 discussion dmitry kononov: what is a reason to initially use the solar chemical composition instead of chemical compositions for stars of later spectral type? valery suleimanov: unfortunately, we do not know the chemical composition of the secondary in ss cyg. therefore, we started from the most conservative case, which is the solar chemical composition. jan-uwe ness: the assumed wd mass in the model (1 m�) seems discrepant from the mass given in the introduction (0.55 m�), a factor 2. this suggest that all parameters of the model may be discrepant from reality, so what do we learn? valery suleimanov: sorry, you mixed up the white dwarf mass with the mass of the secondary (0.55 m�). dmitry bisikalo: how can you discriminate between a boundary layer and a pseudo-photosphere or a hot disk halo? 146 http://dx.doi.org/10.1086/166600 http://dx.doi.org/10.1134/1.1562213 http://dx.doi.org/10.1086/380758 http://dx.doi.org/10.1086/163187 http://dx.doi.org/10.1086/163188 http://dx.doi.org/10.1093/mnras/187.4.777 http://dx.doi.org/10.1088/0004-637x/717/1/363 http://dx.doi.org/10.1086/181378 model atmosphere spectrum fit to the soft x-ray outburst spectrum of ss cyg valery suleimanov: the observed soft xray luminosity is a discrimination factor. any pseudophotosphere or an optically thick hot disk halo with this low surface gravity (log g ≈ 6.2) and the derived effective temperature (≈ 190 kk) would have a radius that is tens of times larger than the white dwarf radius, and the luminosity would be hundreds of times larger than the observed luminosity. 147 introduction model atmospheres results discussion and conclusion 269 acta polytechnica ctu proceedings 2(1): 269–272, 2015 269 doi: 10.14311/app.2015.02.0269 x-ray observations of vy scl-type nova-like binaries in the high and low state p. zemko1,2, m. orio2,3 1sternberg astronomical institute, moscow state university, 13, universitetskij ave., moscow, 119991, russia 2inaf osservatorio di padova, vicolo dell’ osservatorio 5, i-35122 padova, italy 3department of astronomy, university of wisconsin, 475 n. charter str., madison, wi 53704, usa corresponding author: polina.zemko@gmail.com abstract four vy scl-type nova-like systems were observed in x-rays both during the low and the high optical states. they are bz cam, mv lyr, tt ari, and v794 aql. using archival rosat, swift and suzaku observations we found that the x-ray flux for bz cam is higher during the low state, but there is no supersoft x-ray source (sss) that would indicate the thermonuclear burning predicted in a previous article. the x-ray flux is lower by a factor 2–10 in the low than the high state in other systems, and does not reflect the drop in ṁ inferred from optical and uv data. the best fit model for the x-ray spectra is a collisionally ionized plasma model. the x-ray flux may originate in a shocked wind or in accretion onto polar caps in intermediate polar systems that continues even during the low state. keywords: cataclysmic variables dwarf novae x-rays. 1 introduction nova-like (nls) stars are non-eruptive cvs (warner, 1995), classified into several subtypes according to their properties. here we will focus on the vy scl-type nls or “anti-dwarf novae” characterized by the presence of occasional dips on the ilight icurve, iso-called ilow states. the large optical and uv luminosity has suggested that in the high state these objects are undergoing mass transfer onto the wd at a high rate ṁ, > 10−10 m� year−1, sustaining an accretion disk in a stable hot state and preventing dwarf novae (dne) outbursts. the low states have been attributed to a sudden drop of the ṁ from the secondary, or even to a total cessation of a mass transfer (hessman, 2000). greiner et al. (1999) proposed a link between the vy scl-type stars and super soft x-ray sources based on a rosat observations of v751 cyg. these authors found an anti-correlation in the optical and x-ray intensity, and despite the very poor spectral resolution of the rosat hri, the spectrum appeared to be very soft in the low state. these authors suggested that quasi-stable thermonuclear burning occurs on the surface of the wd in the low state. in this framework, vy scl-type stars are key objects in the evolution of interacting wd binaries, in which hydrogen burning occurs periodically without outbursts. thus they may reach the chandrasekhar mass and the conditions for type ia supernovae explosion. using archival x-ray observations, in this paper we compare the high and low state x-ray data of 4 vy scl-type stars attempting to reveal evidence of nuclear burning during the low state, or seeking alternative explanations for the changes that take place during the transition from the high to low state. 2 observation and data analysis we examined the archival x-ray data of vy scl-type stars obtained with swift and rosat and chose the objects that were observed both in the high and low states: bz cam, mv lyr, tt ari and v794 aql. to determine when the low and high optical states occurred, we relied on the data of the variable star network (vsnet) collaboration (kato et al., 2004), aavso1 and asas databases. we used heasoft version 6-13 to analyze the data and xspec version 12.8.0 for spectral modeling. we also estimated the uv magnitudes of the objects using both swift/uvot and additional galex archival observations. 3 results in figure 1 we show the high and low state x-ray spectra of tt ari, bz cam, mv lyr and v794 aql and 1 http://www.aavso.org 269 http://dx.doi.org/10.14311/app.2015.02.0269 p. zemko, m. orio the fits to these spectra. the fit parameters are summarized in table 1 and 2. tt arietis (top panel of fig. 1) is one of the most optically luminous cvs, usually at mag. 10–11. sometimes it abruptly falls into an “intermediate state” around 14 mag. or even into a “low state” at about 16.5 mag. according to belyakov et al. (2010), this binary system consists of a 0.57–1.2 m� tt ari 1 2 3 4 5 6 7 0 .0 0 .1 0 .2 0 .3 bz cam 1 2 3 4 5 6 7 0 .0 0 0 .0 2 0 .0 4 0 .0 6 mv lyr n o rm a liz e d c o u n ts ( cn ts /s /k e v ) 2 4 6 8 0 .0 0 0 .0 2 0 .0 4 v794 aql 1 2 3 4 5 6 0 .0 0 0 .0 4 0 .0 8 0 .1 2 energy (kev) figure 1: the low and high states x-ray spectra of bz cam and mv lyr observed with swift. tt ari and v794 aql were observed with swift during the low and intermediate states. the high and intermediate state spectra are plotted in red and the low state spectra in black. the solid lines show the spectral fits, the dots with error bars are the data points. white dwarf and a 0.18–0.38 m� secondary component. the orbital period is porb = 0.13755114 d. (thorstensen, smak & hessman, 1985). the high state x-ray flux appeared to be about ten times larger than in the low state. we fitted the high state spectrum with a two components thermal plasma model (apec model – emission spectrum from collisionally ionized diffuse gas calculated using the atomdb code v2.0.1), and also with one component black body model. the low state spectrum is also well fitted with a 2 components thermal plasma model. bz cam is a nl star at a distance 830 ± 160 pc. (ringwald & naylor, 1998) with an orbital period of porb = 0.153693(7) d. (patterson et al., 1998). most of the time it shows brightness variations around 12– 13 mag., with rare occasional transitions to the 14– 14.5 mag. low states. bz cam is surrounded by a bright emission nebula with a bow-shock structure. hollis (1992) proposed that the bow shock structure must be due to the interactions of wind from bz cam with the interstellar medium. greiner et al. (2001) suggested that this nebula is photoionized by a bright central object that must be a super soft x-ray source, while the bow shock structure is due to the high proper motion of bz cam. from the second plot of fig. 1 it can be seen that the luminosity is higher in the low state, however, in the very soft spectral region, at energy ≤0.5 kev, the x-ray flux is almost twice higher in the high state, which is exactly the opposite of the idea proposed by greiner et al (1999). interestingly, the spectral fits in both states indicate that we may be observing a strong ne x lyα line at 1.02 kev. we fit the low and high state spectra of bz cam, with a two-component apec model with variable abundances. in this fit we found an underabundance (with respect to solar values) of c and n and and an overabundance of o, ne, and intermediate mass elements like s and ca. figure 2: the swift uvot image of bz cam, in which the bow-shock emission nebula is clearly detected. the uv magnitudes of bz cam in the high and low states indicate a smaller variation than observed in the other objects (albeit in different uv filters). this is explained by figure 2 in which we show the uv image of the nebula obtained with swift/uvot observations. obviously the ionized nebula also emits copious uv flux. comparing the image of bz cam in figure 2 and the one, presented in the figure 4 in (greiner et 270 x-ray observations of vy scl-type nova-like binaries in the high and low state al., 2001), one can see an additional prominent feature of the nebula in the former image. mv lyr is at about 12–13 mag. in the high state and 16–18 mag. in the low state. hoard et al. (2004) have shown that the distance to the object is 505 ± 50 pc. with their fuse (far ultraviolet spectroscopic explorer) observations these authors estimated an upper limit to ṁ during the low state ≤ 3×10−13 m� year−1. the orbital period of this system is porb = 0.1329 d. (skillman, patterson & thorstensen, 1995). our data show that the high state x-ray flux of mv lyr is higher by an order of magnitude than in the low state. the spectrum is also harder with additional component prominent above 1.7 kev. we fitted the high state spectrum of mv lyr with a two components thermal plasma model. a good fit is also obtained with a thermal plasma and a power law model (see table 1). v794 aql varies between 14 and 15 mag. in the high optical state, and in the low states it can plunge to 18 – 20 mag. the orbital period is porb = 0.1533 d. (honeycutt & schlegel, 1985). godon et al (2007) derived the following binary system parameters: mwd = 0.9 m�, high state ṁ = 10−8.5 − 10−8.0 m� year−1, inclination i = 60o, and distance to the system d = 690 pc. we fitted the intermediate state spectrum of v794 aql with two vapec components (see table 2). in both components we need high abundance of mg and ni and underabundant s and fe. 4 discussion and conclusions the first conclusion that can be derived from the presented spectra is that in all the objects both in high and low states there is no evidence of the thermonuclear burning on the surface of wd. an important motivation for this research has been the claim by greiner (1998) and greiner et al. (2001) that some of the wd table 1: fitting models and parameters for bz cam and mv lyr bz cam mv lyr high state low state high state low state satellite swift swift swift swift models vapec vapec+vapec vapec+vapec vapec+plow apec+apec apec+apec n(h)(1022) 0.14 ± 0.03 0.29 ± 0.11 0.6 ± 0.3 0.002+1.049−0.002 0.87 ± 0.23 0.12 +0.34 −0.12 photon index 0.8 ± 0.3 t1 (kev) 10.4 ± 2.7 0.57 ± 0.26 0.6 ± 0.4 0.9 ± 0.7 0.09 ± 0.03 0.7 ± 0.4 t2 (kev) 11 ± 3 63.9+0.1−54.9 12 ± 7 4.8 +42.8 −2.6 flux∗ abs 3.78 ± 0.26 3.95 ± 0.27 5.8+6.4−5.8 5.3 ± 0.8 4.4 ± 0.6 0.9 ± 0.7 flux∗ unabs 4.24 ± 0.26 4.84 ± 0.27 8.6 5.5 ± 0.9 367 ± 50 1.6 ± 1.2 χ2 1.1 1.0 1.0 1.0 1.1 ∗the x-ray flux (×10−12erg cm−2 s−1) was calculated in the following energy range: 0.3 – 10.0 kev for swift xrt table 2: fitting models and parameters for tt ari and v794 aql tt ari v794 aql high state low state intermediate state low state satellite rosat swift swift swift models apec+apec apec+apec apec vapec+vapec apec n(h) (1022) 0.030 ± 0.002 0.056+0.106−0.056 0.026 +0.054 −0.026 0.05 ± 0.04 0.029 +0.086 −0.029 t1 (kev) 0.812 ± 0.014 0.7 ± 0.4 3.6 ± 1.2 16.4+39.1−8.7 8.8 +29.1 −3.7 t2 (kev) 14 ± 9 4.4 ± 1.9 0.97 ± 0.35 flux∗ abs 5.52 ± 0.17 0.97 ± 0.17 0.94 ± 0.15 7.5 ± 1.1 2.1 ± 0.7 flux∗ unabs 6.76 ± 0.17 1.08 ± 0.17 1.02 ± 0.15 8.8 ± 1.1 2.7 ± 0.7 χ2 1.0 1.0 1.2 1.0 0.7 ∗the x-ray flux (×10−12erg cm−2 s−1) was calculated in the following energy ranges: 0.2 – 2.5 kev for rosat pspc and 0.3 – 10.0 kev for swift xrt 271 p. zemko, m. orio in vy scl-type stars must be burning hydrogen quietly in the low state, without ever triggering thermonuclear flashes because of the short duration of the burning. we find that the predicted supersoft x-ray source (manifestation of the thermonuclear burning) does not appear in the low states. we also find another unexpected result, namely that the x-ray luminosity does not follow the optical/uv drop. in one case, bz cam, the x-ray luminosity increases in the low state, and in the others, the decrease in x-ray flux in the low state is smaller than that in the optical and uv. the x-ray flux varies much less than the optical flux, by a factor of 2 to 10 during the transition from high to low state. the best-fit model for the 0.3–10.0 kev broad band spectra is two-component absorbed thermal plasma model. we do not find evidence of a boundary layer of the accretion disk emitting in the x-ray range. probably the emission is mainly in the far uv (e.g godon et al., 2007). it appears likely that some, or all the x-ray flux is produced in a wind from the system: an ongoing fast wind may be the cause of the extended bz cam nebula, initially classified as a planetary nebula. however, tt ari shows a quasi periodic oscillations in x-rays, which can be understood in the context of accretion, but is not explained by a wind. the x-ray flux may also be due to a different, and coexistent mode of accretion other than the disk, i.e. a magnetically driven stream to the polar caps of an intermediate polar. the x-rays and optical flux variations anticorrelate only in bz cam, which may be due to a wind causing additional absorption and obscuring the luminous disk. a second explanation for the lack of correlation of uv/optical versus x-ray flux variations, involving magnetic accretion, appears more plausible at least for three of the systems. the scenario we suggest is that the stream to the polar caps still continues, at decreased rate, when the accretion disks ceases to exist. references [1] belyakov k. v., suleimanov v. f., nikolaeva e. a., borisov n. v., 2010, in werner k., rauch t., eds, american institute of physics conference series vol. 1273 of american institute of physics conference series, p. 342–345. [2] greiner j., 1998, a&a, 336, 626 [3] greiner j., tovmassian g., orio m., lehmann h., chavushyan v., rau a., schwarz r., casalegno r., scholz r., 2001, a&a, 376, 1031 [4] greiner j., tovmassian g. h., di stefano r., prestwich a., gonzález-riestra r., szentasko l., chavarŕıa c., 1999, a&a, 343, 183 [5] godon p., sion e. m., barrett p., szkody p., 2007, apj, 656, 1092 doi:10.1086/510775 [6] hessman f. v., 2000, new astronomy reviews, 44, 155 doi:10.1016/s1387-6473(00)00030-0 [7] hoard d. w., linnell a. p., szkody p., fried r. e., sion e. m., hubeny i., wolfe m. a., 2004, apj, 604, 346 doi:10.1086/381777 [8] hollis j. m., oliversen r. j., wagner r. m., feibelman w. a., 1992, apj, 393, 217 doi:10.1086/171499 [9] honeycutt r. k., robertson j. w., 1998, astronomical journal, 116, 1961 doi:10.1086/300539 [10] honeycutt r. k., schlegel e. m., 1985, pasp, 97, 1189 doi:10.1086/131684 [11] kato t., uemura m., ishioka r., nogami d., kunjaya c., baba h., yamaoka h., 2004, pasj, 56, s1 [12] patterson j., raymond j. c., 1985, apj, 292, 535 doi:10.1086/163187 [13] patterson j., patino r., thorstensen j. r., harvey d., skillman d. r., ringwald f. a., 1996, astronomical journal, 111, 2422 doi:10.1086/117976 [14] ringwald f. a., naylor t., 1998, astronomical journal, 115, 286 [15] skillman d. r., patterson j., thorstensen j. r., 1995, pasp, 107, 545 [16] thorstensen j. r., smak j., hessman f. v., 1985, pasp, 97, 437 [17] warner b., 1995, cataclysmic variable stars (cambridge: cambridge university press). discussion dmitry bisikalo: could you estimate the impact of bow shock to observed x-ray flux? polina zemko: the bz cam image in hα (figure 4 of greiner et al., 2001) shows two distinct bright features in the bow shock front, one at ∼12 arcsec, the other at ∼40-45 arcsec from the central source. nevertheless, the psf of the source in x-ray image obtained with swift is quite symmetric and typical of a point source. so even with the modest spatial resolution of swift we can rule out that the bow shock contributes significantly to the x-ray flux. 272 http://dx.doi.org/10.1086/510775 http://dx.doi.org/10.1016/s1387-6473(00)00030-0 http://dx.doi.org/10.1086/381777 http://dx.doi.org/10.1086/171499 http://dx.doi.org/10.1086/300539 http://dx.doi.org/10.1086/131684 http://dx.doi.org/10.1086/163187 http://dx.doi.org/10.1086/117976 introduction observation and data analysis results discussion and conclusions 90 acta polytechnica ctu proceedings 2(1): 90–93, 2015 90 doi: 10.14311/app.2015.02.0090 simultaneous ubvri observations of the ae aquarii blobs g. latev, r. zamanov, s. boeva, k. stoyanov 1institute of astronomy and national astronomical observatory, bulgarian academy of sciences, 72 tsarighradsko shousse blvd., 1784 sofia, bulgaria corresponding author: glatev@astro.bas.bg abstract we summarize the results of our study of the cataclysmic variable ae aqr on the basis of simultaneous ubv ri observations. for the flares, we estimated the average risetime of about 300 sec, and colours (u − b)0 ∼ 1.1 and (b − v )0 ∼ 0.1. we also calculated temperatures, sizes, masses and expansion velocities of a few individual fireballs. in a single night (16.08.2012), we detected ∼ 8 min quasi-periodic modulation. keywords: stars: individual: ae aqr novae, cataclysmic variables binaries: close stars: flare. 1 introduction ae aquarii is a fascinating magnetic cataclysmic variable (cv) with orbital period of 9.88 h. like the typical cvs, it consists of a k4 iv/v star which transfers material toward a magnetic fast spinning white dwarf (wd). ae aqr has a relatively long orbital period and is one of the largest cvs having semimajor axis a = 2.33 ± 0.02 r�, wd mass mwd = 0.63 ± 0.05 m�, secondary mass m2 = 0.37 ± 0.04 m� (the quantities obtained with high-dispersion time-resolved absorption line spectroscopy by echevarŕıa et al. 2008). a review of ae aqr was presented by p. meintjes (this volume). ae aqr has radio and millimeter synchrotron emission (bastian, dulk & chanmugam 1988), and could be source of tev γ-rays (oruru & meintjes 2012 and references therein). however, magic does not detect γ−ray emission from ae aqr (aleksić et al. 2014). wynn, king & horne (1997) demonstrated that the wd is acting as magnetic propeller and is ejecting most of the matter transferred through the inner lagrangian point in the form of blobs (‘fireballs’). the spitzer infrared spectrum above 12.5 µm can be interpreted as synchrotron emission from electrons accelerated to a power-law distribution in expanding clouds (dubus et al. 2007). ae aqr hosts a rapidly rotating white dwarf. its spin period is p = 33.08 s and spin down rate – 5.64 × 10−14 s s−1 (patterson 1979, mauche et al. 2011), corresponding to a spin-down luminosity of 6 × 1033 erg s−1. a part of this spin-down power goes for ejection of the blobs. the magnetospheric propeller is effective and only a small fraction (∼3%) of the trasferred mateial eventually accretes on to wd (oruru & meintjes 2012). 2 observations our observations of ae aqr were started in 1998 with an electophotometer in v-band and are now extended to simultaneous multicolour (ubv ri) observations with the telescopes of the national astronomical observatory rozhen (2.0m rcc, 50/70 cm schmidt and 60 cm telescope) and the belogradchick astronomical observatory (60 cm telescope). the observational sessions can be summarized as follow: (i) 1998 1999: v-band electrophotometric observations. (ii) aug 2010 aug 2011: simultaneous multicolor (ubv ri) ccd observations. (iii) sep 2011 aug 2013: new simultaneous multicolor ccd observations. observational methods and data reduction are described in zamanov et al. (2012). 3 photometric behavior strong flickering and flaring activity was noted by henize (1949). on time-scales of about 10 minutes, the light curve of ae aqr displays flares with an amplitude up to ≈ 1 mag (see fig.1 and fig.2). multicolour optical photometry was performed by chincarini & walker (1981) in ubv bands. later van paradijs, van amerongen, & kraakman (1989), reported five-colour (in walraven bands) observations and showed that the flares have rise time ∼ 100 − 200 s and occur throughout the whole orbit. in fig.2 we plot the orbital light curve of ae aqr using the orbital period of 0.411655530 d and the zeroorbital phase jd0 = 2449281.4222200 (casares et al. 1996). in this figure a part of our data obtained in august 2013 are shown. the quiescent flux curve is recognizable as smooth orbital variation with two maxima and two minima per orbital cycle. 90 http://dx.doi.org/10.14311/app.2015.02.0090 simultaneous ubvri observations of the ae aquarii blobs figure 1: optical light curve of ae aqr obtained on september 27, 2011 in the bv ri bands. figure 2: the b band light curve of ae aqr (august 2013) folded with the orbital period (9.8797 h). 3.1 the flares as expanding fireballs pearson, horne & skidmore (2003, 2005) formulated analytic expressions for the spectral evolution and continuum light curves of flickering and flaring variability that occur over a wide range of astrophysical objects. they applyed these expressions to the observations of the cataclysmic variables ae aqr and ss cygni, and of the supernova sn 1987a, deriving physical parameters for the material involved. they have shown that the observed flare spectrum and evolution of ae aqr is reproducible with an isothermal fireball with population ii abundances. the interested reader is directed to their papers for a full discussion. the basic assumptions are: (i) the flares of ae aqr are due to the appearance and expansion of fireballs (blobs); (ii) the blobs are isothermal; (iii) they represent spherically symmetric expansion of a gaussian density profile with radial velocity proportional to the distance from the center of the expansion. figure 3: upper panel: dereddened fluxes of a fireball at the maximum of the flare. the solid line is a black body fit. lower panel: time evolution of the flare in the v band. the (red) solid line represents the isothermal fireball model. following pearson et al. (2005), the dimensionless time β is defined as β = 1 + h(t − tpk), where tpk is the time of the peak of the flare, h is an “expansion constant” setting the speed of the expansion. the dimensionless time β is also the expansion factor being the constant of proportionality between the current and peak scale length apk: β ≡ a/apk. the central density of the fireball is ρ = m (πa2)3/2 , (1) where m is the total mass of the material involved in the expansion (the fireball mass). the speed of expansion at a is v = ha. the optical depth parallel to the 91 g. latev et al. observer’s line of sight is τ(y) = − ∫ −∞ ∞ κ dx = τ0 e −2( y a )2 (2) where y is the impact parameter (the distance from the fireball center perpendicular to the line of sight), κ1 is the linear absorption coefficient, � is the correction for stimulated emission, � = 1 − e(hν/kt). here t is the fireball temperature, a is the length scale (which we call the fireball size). the optical depth on the line of sight through the center of the fireball (y = 0) is τ0 = κ1 � m 2 21/2 t 1/2 0 ν 3 π5/2 a5 . (3) the emission of the fireball is: fν = π a2 bν(t) 2 d2 s(τ0), (4) where s(τ0) is the ”saturation function” (see fig. 1 of pearson et al. 2005). to calculate the fireball parameters t,m,a,h,ρ and v, we performed the following: 1. we compute the peak flux of the fireball, fpk, in the five optical bands (ubv ri). an example is given in fig.3, where the calculated peak fluxes (corrected for the reddening) are plotted. 2. we derive the temperature of the fireball with a black body approximation applying iraf nfit1d routine. 3. we evaluate the size of the blob at the peak, apk, using τ0 = 6.8202 and eq.4. 4. we calculate the mass of the fireball (eq.3). the calculations are done adopting population ii abundances (κ1 = 1.27 × 1052 m−1). 5. fitting the v band light curves, we derive the expansion constant h. an example is given in fig.3 lower panel) . 6. we derive the speed of expansion and the central density. part of the calculated parameters are given in table 1. we reach lovely agreement between between the model and the light curves of the optical flares using fireballs with a temperature t ∼ 15000 k and mass m ∼ 1020 g. blobs are detected in two other close binaries containing white dwarfs (the recurrent nova rs oph and the symbiotic star ch cyg). in future it will be interting to understand is it the same mechanism (magnetic propeller), which generates the blobs. figure 4: a) light curve (in johnson b band) of ae aqr obtained on august 16, 2012. quasi-periodic oscillations are visible. b) power spectrum of the light curve. the maximum indicates a period of about 8 min. 4 unusual behaviour on august 16, 2012 usually, ae aqr has light curve as shown in fig.1. however, in our observation obtained on august 16, 2012 (see fig.4) a clear periodicity is visible. the power spectrum has a maximum corresponding to t = 7.5 ± 0.2 min. possible explanations of this unusual behavior are: (1) beat modulation between the orbital period and the wd spin period (e.g. as supposed for bg cmi, fig.7.3 of warner 1995). the equation 1/pbeat = 1/pspin − 1/porb (5) gives pbeat = 33.03 s, which is too short in comparison with the observed value. (2) blob rotating with keplerian velocity at the border of the magnetosphere: rblob = (2π) −2/3 (gmwd) 1/3 t2/3, (6) the period t = 7.5 min corresponds to keplerian rotation at distance rblob ≈ 0.11 r�. the magnetospheric radius is estimated as rm ≈ 4−8rwd ≈ 0.05−0.10 r� (zamanov et al. 2012 and references therein). in our opinion, we had the chance to observe a blob rotating at the magnetosphere boundary. 92 simultaneous ubvri observations of the ae aquarii blobs table 1: the computed parameters of the fireballs. the time is in format yyyymmdd hh:mm. in the table are given as follows: rise time in seconds, colours of the peak emission of the fireball (corrected for the interstellar reddening), temperature of the fireball, its mass and size (at the peak of the flare). quantity 20100813 23:40 20100814 19:20 20100814 19:48 20100814 20:10 20110831 21:43 rise time [sec] 260±20 230±25 290±30 440±20 260±30 (u − b)0 -1.36±0.06 -1.43±0.03 -0.93±0.04 -0.80±0.05 -1.02±0.07 (b − v )0 0.24±0.03 0.23±0.03 0.17±0.03 0.19±0.03 0.03±0.06 temperature t [k] 14 545±1000 27 292±1500 10 856±150 9 527±100 13 395±200 mass m [1019 g] 9.6±1.5 6.8±1.5 39 ±6 78±12 97±15 size apk [10 9 cm] 3.0±0.3 2.5±0.3 5.3±0.3 7.1±0.4 7.7±0.4 5 conclusions using 4 telescopes, we performed simultaneous observations in 5 bands (ubv ri) of the flare activity of the cataclysmic variable ae aqr. adopting the model of an isothermically expanding ball of gas, we calculated parameters (temperature, size, mass) of a few individual fireballs. in a single night, we detected ∼ 8 min quasiperiodic modulation, which might be due to a blob rotating at the magnetosphere boundary. in future we intent to measure more blobs and to search for correlations between their physical parameters. also, we will try to do a follow-up study of the qpos phenomenon. acknowledgement the authors are very grateful to the anonymous referee for useful notes and comments. this work was supported in part by the op ”hrd”, esf and the bulgarian ministry of education and science (bg051po0013.3.06-0047). references [1] aleksić, j., ansoldi, s., antonelli, l. a., et al. 2014, a&a, 568, aa109 [2] bastian, t. s., dulk, g. a., & chanmugam, g. 1988, apj, 324, 431 doi:10.1093/mnras/282.1.182 [3] casares, j., mouchet, m., martinez-pais, i. g., & harlaftis, e. t. 1996, mnras, 282, 182 [4] chincarini, g., & walker, m. f. 1981, a&a, 104, 24 doi:10.1111/j.1365-2966.2008.13248.x [5] dubus, g., taam, r. e., hull, c., watson, d. m., & mauerhan, j. c. 2007, apj, 663, 516 [6] echevarŕıa, j., smith, r. c., costero, r., zharikov, s., & michel, r. 2008, mnras, 387, 1563 doi:10.1111/j.1365-2966.2012.20410.x [7] henize, k. g. 1949, aj, 54, 89 [8] oruru, b., & meintjes, p. j. 2012, mnras, 421, 1557 doi:10.1046/j.1365-8711.2003.06079.x [9] patterson, j. 1979, apj, 234, 978 doi:10.1086/426582 [10] pearson, k. j., horne, k., & skidmore, w. 2003, mnras, 338, 1067 [11] pearson, k. j., horne, k., & skidmore, w. 2005, apj, 619, 999 [12] van paradijs, j., van amerongen, s., & kraakman, h. 1989, a&as, 79, 205 [13] warner, b. 1995, cambridge astrophysics series, 28 doi:10.1002/asna.201211718 [14] wynn, g. a., king, a. r., & horne, k. 1997, mnras, 286, 436 [15] zamanov , r. k., latev, g. y., stoyanov, k. a., et al. 2012, astronomische nachrichten, 333, 736 discussion ashley pagnotta: what is the timescale of the short unexplained flares superimposed on the orbital modulations? georgi latev: sampling every 10 seconds, average time from peak to peak is about 8 minutes simone scaringi: given the simultaneous multi-color photometry can anything be said about color evolution of ae aqr? g. latev: we will try to do this in our future work. 93 http://dx.doi.org/10.1093/mnras/282.1.182 http://dx.doi.org/10.1111/j.1365-2966.2008.13248.x http://dx.doi.org/10.1111/j.1365-2966.2012.20410.x http://dx.doi.org/10.1046/j.1365-8711.2003.06079.x http://dx.doi.org/10.1086/426582 http://dx.doi.org/10.1002/asna.201211718 introduction observations photometric behavior the flares as expanding fireballs unusual behaviour on august 16, 2012 conclusions 127 acta polytechnica ctu proceedings 1(1): 127–131, 2014 127 doi: 10.14311/app.2014.01.0127 multifrequency behaviour of the gamma-ray binary system psr b1259-63: modelling the fermi flare brian van soelen1, pieter j. meintjes1 1department of physics, university of the free state, bloemfontein 9300, south africa corresponding author: vansoelenb@ufs.ac.za abstract this paper presents a brief overview of the multifrequency properties of the gamma-ray binary system psr b1259-63 from radio to very high energy gamma-rays. a summary is also presented of the various models put forward to explain the fermi “flare” detected in 2011. initial results are presented of a new turbulence driven model to explain the gev observations. keywords: gamma-ray binaries individual: psr b1259-63. 1 introduction there are only a handful of known gamma-ray binary systems. of these, it is only in psr b1259-63 that the nature of the compact object is known due to the detection of pulsed radio emission (johnston et al., 1992a). the very high energy (vhe) nature of the source was confirmed by the h.e.s.s. telescope around the 2004 periastron passage (aharonian et al., 2005). the gammaray binary system consists of a 48 ms pulsar, in an eccentric (e = 0.87) 3.4 year orbit around a be star (e.g. johnston et al., 1992, 1994), with a spin-down luminosity of ė = 8.3 × 1035 erg s−1 (e.g. wang et al., 2004). recent spectroscopic observations have placed better constraints on the parameters of the optical companion, ls 2883, and the binary system (negueruela et al., 2011). in this paper, we present an overview of the multifrequency behaviour of psr b1259-63, and discuss some of the modelling undertaken to explain the large “flare” event detected by fermi after the previous periastron passage. 2 multifrequency observations in this section we summarize the multifrequency behaviour of the non-thermal emission originating from psr b1295-63, discussing the radio, x-ray, vhe gamma-ray and, finally, the gev gamma-ray observations. 2.1 radio behaviour psr b1259-63 was originally detected as a radio pulsar during a galactic plane survey undertaken with the 64 m parkes radio telescope (johnston et al., 1992b). observations around periastron showed that there was an increase (and variability) in the dispersion measurement of the pulsed signal and an eclipse of the pulsed signal between approximately 20 days before periastron to 20 days after (e.g. johnston et al., 1996, 1999, 2001). significantly, around periastron an unpulsed and variable radio emission was detected with a flux density of 10 − 30 mjy, higher than the flux of the pulsed signal (2 − 3 mjy). these observations are explained by the pulsar passing through and/or behind the circumstellar disc of the be star, and the formation of synchrotron emission in the pulsar wind nebula. more recently a large extended pulsar wind nebula has also been reported around periastron by moldón et al., (2011). 2.2 x-ray behaviour a non-thermal x-ray component has been detected from psr b1259-63 across the whole of the orbit, with early detections determining a luminosity of l ∼ 1034 erg s−1 around periastron (kaspi et al., 1995). combined asca, chandra, xmm -newton, bepposax, swift and suzaku observations covering the 1997 – 2010 periastron passages have shown a consistent x-ray light curve (see chernyakova et al., 2006, 2009, abdo et al., 2011, and references therein). these datasets have presented particular coverage focused on the disc crossing epochs. the observations show that the x-ray emission peaks within two narrow time bands most likely associated with the pulsar approaching the circumstellar disc. pavlov et al., (2011) also reported on an extended x-ray structure detected by chandra near apastron. the observations in the 0.5–8 kev en127 http://dx.doi.org/10.14311/app.2014.01.0127 brian van soelen, pieter j. meintjes ergy band showed an elongated structure (∼ 4′′) with a luminosity of ∼ 1032 erg s−1 (assuming a distance of 3 kpc). it is likely that this is associated with the pulsar wind nebula outflow. 2.3 vhe behaviour psr b1259-63 was detected for the first time at vhe with the h.e.s.s. telescope during the 2004 periastron passage (aharonian et al., 2005). the observations showed a variable light curve with an implied maximum flux a few days before and after periastron. the system has since been detected during 2007 (aharonian et al., 2009) and 2011 (abramowski et al., 2013), albeit with much less coverage. the h.e.s.s. observations appear to be consistent with the double peak structure observed at x-ray and radio wavelengths, but this conclusion is not definitive due to lower coverage during the pre-periastron approach of the pulsar. the spectrum of the vhe emission also appears to be consistent with a single power-law distribution, with no cut-off in the spectrum (see e.g. fig. 4 in abramowski et al., 2013). while the coverage over the last periastron passage was poor, the 2011 observations were coincident with the fermi “flare” discussed in the next section. unlike the fermi detection, the h.e.s.s. observations showed no significant change between the preand post-flare event, and morphologically the gev to tev energies are not consistent with a single power-law distribution (see fig. 5 in abramowski et al., 2013). 2.4 he behaviour due to the consistency of the radio, x-ray and vhe light curves, it was presumed (before 2010/2011) that fermi would observe similar behaviour. in fact, it was not actually clear whether the system would be detectable by fermi (e.g. chernyakova et al., 2009). the gamma-ray detection and rapid transient behaviour that were reported by fermi were completely unexpected. as reported by abdo et al. (2011), the system was relatively undetected before periastron, with a faint “brightening” phase which began around the point of the first disc crossing and lasted until ∼ 18 days after periastron. however, approximately 30 days after periastron, there was a rapid increase in the detected emission from the source showing an average flux of f(> 100mev) = (4.4±0.3stat ±0.7sys)×10−10 erg cm−2 s−1, with a photon index of γ = (1.4 ± 0.6stat ± 0.2sys) and a cutoff energy ec = (0.3 ± 0.1stat±sys) gev. this “flare” was detected until > 60 days after periastron. significantly, radio, x-ray, and vhe observations showed no unusual or flaring events. most surprisingly, the daily binned light curve showed a peak luminosity of lγ ≈ 8 × 1035 erg s−1 (for isotropic emission at an assumed distance of 2.3 kpc), corresponding to ≈ 100 per cent of the spin-down power of the pulsar. the emission detected by fermi has proved difficult to reconcile with models of the system. in the next section we discuss some of the models pertaining to gev energies and the fermi flare. 3 modelling the fermi flare 3.1 effect of the infrared excess in anticipation of the fermi observations of psr b125963, van soelen & meintjes (2011), van soelen et al. (2012) and meintjes & van soelen (2012), considered the influence the infrared flux from the circumstellar disc of the be star would have on gamma-ray production. if the same population of electrons which produces the vhe emission through inverse compton scattering of the optical photons, scatters the infrared photons from the disc, it will produce emission in the gev energy range. first using an isotropic approximation (van soelen & meintjes, 2011) and then a full anisotropic approximation, which took into account the anisotropic inverse compton scattering as well as the changing flux from the circumstellar disc as measured from the pulsar’s position (van soelen et al., 2012), we showed that the scattering of the infrared photons would increased the flux at gev energies by a factor of 2. while it was thought that the large size of the circumstellar disc would result in it dominating near the disc crossing, we found that the low density of the disc mitigated this effect as the majority of the infrared emission occurred near the star and not far out in the disc. however, the flux was below what was detected by fermi. 3.2 cold pulsar wind khangulyan et al., (2012) proposed that the scattering of the cold pulsar wind (γ = 104) off photons from the circumstellar disc, after the second disc passage, could produce the observed fermi flare. the authors propose that the shock-front around the pulsar is confined within the disc, and after the pulsar leaves the circumstellar disc the opening of the shock structure will allow the inverse compton scattering of the cold pulsar wind to produce the flare. one problem presented by this model is the number of additional photons required to do this. the authors estimate that the additional photon distribution must have a luminosity of approximately 40 per cent of the stellar luminosity. this does not appear to be consistent with the observations obtained around periastron nor with the estimation of the available energy from shock-heating (van soelen et al., 2012). however, it must be cautioned that the reported observations were not undertaken at the same time as 128 multifrequency behaviour of the gamma-ray binary system psr b1259-63: modelling the fermi flare the fermi flare and new observations around the next periastron passage must be obtained, to obtain better constraints. 3.3 sph modelling a process of smooth particle hydrodynamic (sph) modelling of psr b1259-63 was undertaken by takata et al., (2012) to model the behaviour of the be star’s circumstellar disc around periastron. the simulations show that the disc should be considerably disrupted during the pulsar crossing. however, observations on 5 january 2011, approximately three weeks after periastron, showed that the mid-infrared flux was consistent with previous observations (see van soelen et al., 2012). this suggests that either the disc was not dramatically effected, or that there was a rapid recovery. in addition, the approximation of the emission light curve, while in agreement with the x-ray emission, was not in agreement with the gev light curve, as the simulations suggested that the maximum emission should occur before periastron (see figs. 3 and 5 in takata et al., 2012). 3.4 doppler boosting emission the most recent model which has been suggested is that the synchrotron emission is doppler boosting to gev energies to produce the observed flare. this was modelled in detail by kong et al. (2012). the authors consider two emission regions along the shock front: the first near the apex with a bulk lorentz factor of γ = 1 and the second along the shock tail which is mildly relativistic (γmax ≈ 2) and is beamed along the shock. these simulations are able to account for the fermi detection during the flare period, but are not so successful during the pre-periastron period – though it should be noted that the pre-periastron detections have a much lower significance. the light curve predictions also suggest only a single peaked emission around periastron for the gev energies, and a double peaked structure from the x-ray and vhe emission. however, the modelled light curves predict a higher flux for the gev and tev emission around periastron (see fig. 5 in kong et al., 2012), an effect which is not observed. 3.5 turbulence driven emission? thus far all models have presented unique problems in modelling the fermi flare. for this reason it is still necessary to consider alternative models. below we consider the possibility of a turbulence induced flare. hydrodynamical simulations of colliding wind systems have shown that kelvin-helmoltz instabilities develop along the dividing line between the pulsar and stellar winds (see e.g. bosch-ramon et al., 2012). if these instabilities form a turbulent medium, it may be possible for secondary particle acceleration to occur due to second-order fermi acceleration (see e.g. bosch-ramon & rieger 2012). it has been considered by a number of authors that under the effect of radiative cooling and second-order fermi acceleration, a maxwellian-like distribution will form (see e.g. schlickeiser 1985 for a discussion of firstand second-order acceleration with radiative cooling). such a secondary distribution of electrons may be responsible for producing the gev emission in psr b1259-63 if acceleration in the turbulent region occurs around the second disc crossing. as a first approximation we consider the modified maxwellian distribution suggested by stawarz & petrosian (2008), given by ne = n0γ 2 exp [ − 1 a ( γ γeq )a] , (1) where a = 3 − q for dominant synchrotron/thomsonlimit inverse compton cooling, or a = 1.5 − q for dominant klein-nishina inverse compton cooling. here q is the index of the turbulence power-law, and γeq is determined by equating the appropriate radiative cooling time to the acceleration time (see stawarz & petrosian, 2008, and reference therein, for details related to the acceleration). we calculate the best fit inverse compton emission for the electron distribution given by equation (1) for isotropic scattering (e.g. blumenthal & gould, 1970), with a blackbody distribution of photons (t? = 33 000 k). in this first approximation an isotropic radiation field has been assumed. the equilibrium energies required to fit the maxwellian distributions (equation 1) were found to be γeq = 1614.4 for the bohr approximation (q = 1), γeq = 656.30 for the hard-sphere approximation (q = 2), γeq = 1030.4 for the kolmogorov spectrum (q = 5/3) and γeq = 1196.7 for the kraichnan spectrum (q = 3/2). only a = 3 − q was considered. a constant factor was applied to match the flux level of the flare. the resulting fits are shown in fig. 1. the figure shows that the distribution produces a reasonable fit to the resulting fermi detection. initial considerations of the parameter space suggest that it may be possible to accelerated electrons to the appropriate energies around the period of periastron. a more detailed analysis is currently under-way. 129 brian van soelen, pieter j. meintjes -12 -11 -10 7 8 9 10 11 lo g ν f ν [e rg s -1 c m -2 ] log energy [ev] q=2 q=1 q=5/3 q=3/2 figure 1: inverse compton spectrum formed by a modified maxwellian electron distribution fitted to the post-periastron fermi flare detection. references [1] abdo, a.a., ackermann, m., ajello., m. et al.: 2011, apj, 736, l11 doi:10.1088/2041-8205/736/1/l11 [2] aharonian, f., et al. (h.e.s.s. collaboration): 2005, a&a, 442, 1 [3] aharonian, f., et al. (h.e.s.s. collaboration): 2009, a&a, 507, 389 [4] abramowski, a., et al. (h.e.s.s. collaboration): 2013, a&a, 551, a94 [5] blumenthal, g.r., gould, r.j.: 1970, reviews of modern physics, 42, 237 [6] bosch-ramon, v., rieger f.m.: 2012, in astroparticle, particle, space physics and detectors for physics applications proceedings of the 13th icatpp conference.. g. simone et al. (eds.), 219 [7] bosch-ramon, v., barkov, m.v., khangulyan, d., et al.: 2012, a&a, 544, 59 [8] chernyakova, m., neronov a., lutovinov a., et al.: 2006, mnras, 367, 1201 doi:10.1111/j.1365-2966.2005.10039.x [9] chernyakova, m., neronov, a., aharonian f., et al.: 2009, mnras, 397, 2123 doi:10.1111/j.1365-2966.2009.15116.x [10] johnston, s., manchester, r.n., lyne, a.g, et al.: 1992a, apj, 387, 37 [11] johnston, s., lyne, a.g, manchester, r.n., et al.: 1992b, mnras, 255, 401 [12] johnston, s., manchester, r.n., lyne, a.g, et al.: 1994, mnras, 268, 430 [13] johnston, s., manchester, r.n., lyne, a.g, et al.: 1996, mnras, 279, 1026 [14] johnston, s., manchester, r.n., mcconnel, d.: 1999, mnras, 302, 277 [15] johnston, s., wex, n., nicastro, l., et al.: 2001, mnras, 326, 643 [16] kaspi, v. m., tavani m., nagase f., et al.: 1995, apj, 453, 424 doi:10.1086/176403 [17] khangulyan, d., aharonian, f.a., bogovalov, s.v., et al.: 2012, apj, 752, l17 doi:10.1088/2041-8205/752/1/l17 [18] kong, s. w., cheng, k. s., huang, y. f.: 2012, apj, 753, 127 doi:10.1088/0004-637x/753/2/127 [19] meintjes, p.j., van soelen, b.: 2012, in multifrequency behaviour of high energy cosmic sources, f. giovannelli, l. sabau-graziati (eds.) mem. s.a.it., 83, 246 [20] moldón j., johnston s., ribó m., et al.: 2011, apj, 732, l10 [21] negueruela i., ribó m., herrero a., et al.: 2011, apj, 732, l11 [22] pavlov g. g., chang c., kargaltsev o.: 2011, apj, 730, 2 [23] schlickeiser, r.: 1985, a&a, 143, 431 [24] stawarz, l., petrosian, v.: 2008, 681, 1725 [25] takata, j., okazaki, a. t., nagataki, s. et al.: 2012, apj, 750, 70 [26] van soelen, b., meintjes, p.j.: 2011, mnras, 412, 1721 [27] van soelen, b., meintjes, p.j., odendaal, a., et al.: 2012, mnras, 426, 3135 [28] wang, n., johnston, s., manchester, r. n.: 2004, mnras, 351, 599 discussion carlotta pittori: in august 2010 agile reported the detection of gamma-ray activity above 100 mev from the psr b1259-63 region during the initial approach-to-periastron part of the orbit. fermi did not confirm this detection, but later in november reported faint gamma-ray emission before the big 2011 flare. are you aware of the agile detection and can the model explain a ∼ 100 mev emission just before the disk passage? 130 http://dx.doi.org/10.1088/2041-8205/736/1/l11 http://dx.doi.org/10.1111/j.1365-2966.2005.10039.x http://dx.doi.org/10.1111/j.1365-2966.2009.15116.x http://dx.doi.org/10.1086/176403 http://dx.doi.org/10.1088/2041-8205/752/1/l17 http://dx.doi.org/10.1088/0004-637x/753/2/127 multifrequency behaviour of the gamma-ray binary system psr b1259-63: modelling the fermi flare brian van soelen: yes i am aware of the reported agile detection, as well as the fermi follow-up analysis that did not detect emission. at this point i do not believe any model has completely accurately modelled the events that occurred around the previous periastron passage, including the reported detection by agile. 131 introduction multifrequency observations radio behaviour x-ray behaviour vhe behaviour he behaviour modelling the fermi flare effect of the infrared excess cold pulsar wind sph modelling doppler boosting emission turbulence driven emission? acta polytechnica ctu proceedings doi:10.14311/app.2016.3.0001 acta polytechnica ctu proceedings 3:1–6, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app statistical evaluation of fatigue data of components chi nghia chung∗, zoltan major johannes kepler university linz, institute of polymer product engineering, altenbergerstraße 69, 4040 linz, austria ∗ corresponding author: chi_nghia.chung@jku.at abstract. a variety of steels, cast iron grades and other metals have long been used for the production of machine components. in recent years, however, new materials such as sintered materials and plastics become increasingly important. because of the large number of different fibers, matrices, stacking sequences, processing conditions and processes and the variety of resulting material configurations it is not possible to rely on proven fatigue models for conventional materials. moreover, the development of models, which are valid for all composites are generally extremely difficult. in this work, a possible application of high-performance composites as materials for machine elements are investigated. this study attempts to predict the fatigue behavior and the consequent durability, based on laboratory measurements. using the statistics program jmp, the aquired data was subjected to a reliability analysis in order to ensure the plausibility, validity and accuracy of the measured values. keywords: fatigue models, statistical evaluation, reliability, techniques of parameter fitting, jmp. 1. introduction there was a tremendous advance in the field of plastics took in the past few decades. plastics have become an integral part of our daily life. polymers are flexible materials which can cover a great range of applications. they replace more and more metal, glass, wood and other materials. the development of novel artificial materials often opens a door to new technologies. electrical and automotive industry are hard to imagine without plastics, the use of polymer materials has revolutionized medicine. cyclically stressed components have a limited durability, therefore it is important to perform fatigue tests or simulations on critical components to predict their lifetime. figure 1 shows representative loading patterns with constant amplitudes for the whole loading level [1]. figure 1. constant amplitude loading patterns. fatigue tests are performed to study the relationship between the fatigue resistance of a material, component or structural element and cyclic loading [2]. it is a slowly progressing damage process. the strengths lie far below the static strength and yield strength. the objective of this work was to create high cycle fatigue curves from experimental datasets, using suitable material models, similar to the wöhler curve. the curve should as closely as possible reflect the experimental values. since specimens and test conditions are never 100 % identical, in two discrete measurements there is always scattering in the results, which can span over a decade in fiber reinforced polymers [1]. therefore to handle and correctly interpret experimental results, statistical methods must be used. 2. fatigue models in general, fatigue models are quantifications of physical material properties. they are independent of the shape of the component and are usually based on experimentally acquired data. the aim of all models is to predict how a component behaves in certain conditions. the relationships are reproduced mathematically afterwards, therefore they are mathematical models. in this work, the focus will be on presenting fatigue models [3]. in order to design a component correctly in terms of fatigue, a complete set of experimentally acquired data is usually required. since this is not possible for reasons of time and cost, engineers have to rely on predictive models. this models predict the durability n (in cycles) under a given cyclic loading. 1 http://dx.doi.org/10.14311/app.2016.3.0001 http://ojs.cvut.cz/ojs/index.php/app chi nghia chung, zoltan major acta polytechnica ctu proceedings 2.1. the basquin model the basquin fatigue model is a linear regression model. in a logarithmic scale, the durability (logn) is plotted versus the stress amplitude (log∆σ). log n = a−b · log∆σ, ∆σ ≥ ∆σ0, b ≥ 0 (1) a and b are material parameters that need to be approximated with appropriate fitting methods. in the model, ∆σ is limited by ∆σ0, the endurance limit. in this work, the endurance limit is at 406 cycles. ∆σ0 itself has no direct influence on the model, since it is not considered in the formula. 2.2. the strohmeyer model in contrast to the basquin model the models of strohmeyer and weibull are nonlinear. they are shaped by smoothing a piecewise linear function. again a and b are the material parameters. log n = a−b · log(∆σ − ∆σ0), b ≥ 0 (2) 2.3. the weibull model the weibull model is more complex, since there are more parameters to fit. log(n+d) = a−b·log ( ∆σ − ∆σ0 ∆σst − ∆σ0 ) , b ≥ 0 (3) a, b and d describe the material parameters and ∆σst denotes the ultimate strength. 2.4. the bastenaire model logn −b ∆σ − ∆σ0 = a ·exp[−c · (∆σ − ∆σ0)] (4) a, b and c express the material parameters. this model will be henceforth denoted as bastenaire1, because one can find another model of bastenaire in the literature, which is called bastenaire2 here. it is quite similar to equation 4 and it is also discussed for the purpose of comparison. n = a ·exp[−c · ∆σ − ∆σ0 b ]/(∆σ − ∆σ0) (5) 3. techniques of parameter fitting in this section, two exemplar fitting methods are discussed in greater detail [4–6]. both methods are current and proven estimation methods in statistics. 3.1. the method of least squares (l2-norm) the method of least squares is a mathematical standard procedure for compensation calculation. here a function is determined, which fits a point cloud as close as possible. a point cloud is a scatter plot, a graphical representation of statistical measurements in a coordinate system. to illustrate the method, the basquin model is used as an example. consider a dependent variable y (in this case, logn), which depends on one or several variables (in this case a, b and log∆σ). the relationship between the dependent variable and the arguments are described in a model function f [7]. the model function f can be linear, as in this case, but it can also have any other shape (parabolic, exponential, ...). the general form is: y(x) = f(x; a1.........am) (6) in this case: log n = a−b · log∆σ, ∆σ ≥ ∆σ0, b ≥ 0 (7) the parameters a and b should be adapted such that bad data points (outliers) have only little effects on the fitting. if no unique solution that perfectly fits the point cloud can be found, then a compromise solution with the smallest overall deviation from the point cloud is the valid criterion. for this purpose, the sum of the squares of the differences between the model function f and the measured values yi is formed. n∑ i=1 (f(xi; a1.........am) −yi)2 (8) the parameters a and b are adapted until the sum of squares becomes minimal. in this work the fitting procedure carried out with the help of a solver implemented in excel. 3.2. the method of least absolute deviations (l1-norm) the l1 norm is a more robust fitting method. outliers are not so strongly weighted. in principle, this method works similarly to the previously explained method of least squares. instead of the sum of the squares, the absolute sum of the differences between the model function f and the measured values yi is calculated. abs ‖ n∑ i=1 (f(xi; a1.........am) −yi) ‖ (9) subsequently, the parameters a and b are also be adjusted until the absolute sum is minimized. 4. reliability the reliability is a measure of the accuracy of the measurement, as well as for the reliability of data. measurement series with very high reliability are therefore almost free of random errors, which means that they are repeatable at any time under the same measurement conditions and thereby provide approximately the same results [8, 9]. therefore, one gets a high reliability, when performing controlled and standardized measurements. to examine the reliability, different techniques can be used [10]. some known techniques are: 2 vol. 3/2016 statistical evaluation of fatigue data of components • test retest reliability • parallel forms reliability • split half reliability • internal consistency reliability. reliability analysis can be easily carried out with various computer programs. known programs are spss or, as used in this work, jmp. 5. evaluation of fatigue data two different materials were tested. material#1 is a glass fiber reinforced, semi-crystalline thermoplastic. material#2 is a carbon fiber reinforced, semicrystalline thermoplastic, where the bearing was simulated as a compliant bearing. material#3 is the same material as material#2, however it has been simulated as a rigid bearing. 5.1. fatigue models figure 2 exhibits the applied fatigue models for material#1. since for material#2 the ultimate strength and the load at the endurance limit could not be investigated, only the basquin model could be applied 3. in the case of material#3 the fitting with the strohmeyer and the weibull model resulted in the same curve 4. figure 2. applied fatigue models for material#1. figure 3. applied fatigue models for material#2. figure 4. applied fatigue models for material#3. the software minimized the deviation between the models and the measured values. by comparing the obtained model parameters the optimal model can be easily determined, not only qualitatively but also quantitatively. figure 5 and 6 show the optimal fatigue models for the materials. since for material#2 only the basquin model was applied, it was omitted below. figure 5. optimal fatigue model for material#1. figure 6. optimal fatigue model for material#3. 5.2. reliability analysis in jmp in jmp custom tables can be created, or files of different formats (excel, sas, text files, ...) can be processed. for a correct data input, the appropriate settings in the import wizard must be applied. in order to determine the best distribution for the measured values, a life distribution is performed [11]. the program returns a table with the appropriate distributions for the respective materials. jmp sorts them in descending order, the best model being on the top. the distribution of the measured data is weighted by 3 criteria.the akaike information criterion (aic), the bayesian information criterion (bic) and loglikelihood (maximum probability) used as estimation methods for the selection of models in statistics. 3 chi nghia chung, zoltan major acta polytechnica ctu proceedings tables 1, 2 and 3 show the "model comparisons" for each materials. according to the distribution analysis in jmp, the measured values from material#1 follows a natural logarithmic distribution, from material#2 a frechet distribution and from material#3 a weibull distribution. distribution aicc loglike bic lognormal 503.65435 498.90435 504.79323 weibull 503.66928 498.91928 504.80816 frechet 505.47261 500.72261 506.61149 loglogistic 505.74393 500.99393 506.88280 exponential 518.31862 516.08332 519.02776 lev 538.93104 534.18104 540.06992 logistic 546.20378 541.45378 547.34266 normal 550.30607 545.55607 551.44495 sev 563.21318 558.46318 564.35205 table 1. comparisons of the distribution of the measured data for material#1. distribution aicc loglike bic frechet 297.91571 292.41571 297.21150 lognormal 299.16181 293.66181 298.45760 loglogistic 300.06766 294.56766 299.36345 weibull 301.43592 295.93592 300.73171 exponential 305.28286 302.83841 305.23631 lev 320.31613 314.81613 319.61192 logistic 326.13477 320.63477 325.43056 normal 329.93763 324.43763 329.23342 sev 337.33938 331.83938 336.63517 table 2. comparisons of the distribution of the measured data for material#2. distribution aicc loglike bic weibull 407.25467 402.16376 407.44188 loglogistic 409.33460 404.24369 409.52181 lognormal 409.74094 404.65003 409.92814 frechet 415.33191 410.24100 415.51912 exponential 416.03305 413.69971 416.33877 lev 432.98688 427.89597 433.17409 logistic 439.48273 434.39182 439.66994 normal 439.98404 434.89313 440.17125 sev 445.68650 440.59560 445.87371 table 3. comparisons of the distribution of the measured data for material#3. based on this findings, the durability evaluation was performed. tables 4, 5 and 6 display the mean estimates for each material. σ is the standard deviation from the measured data, β0 and β1 denote positional and shape parameters. stderror stands for standard error and describes the standard deviation, but from the estimate function. additionally the tables show for each estimation the 95 % confidence interval. µ is the estimated average value of cycles, which is dependent on the loading. strictly speaking, this designation should be logµ, because the relations are from logarithmic nature. par. estimate stderror low95 % up95 % β0 79.25810 6.98804 64.83945 93.67675 β1 -14.17640 1.44835 -17.16484 -11.18796 σ 0.92287 0.14971 0.69226 1.31688 table 4. mean estimations for material#1. µ = 79.2581 − 14.1764 · log(loading) (10) par. estimate stderror low95 % up95 % β0 126.9802 20.30212 82.06733 168.4553 β1 -22.7520 3.98476 -30.90587 -13.9531 σ 0.6862 0.15914 0.45366 1.1457 table 5. mean estimations for material#2. µ = 126.9802 − 22.75204 · log(loading) (11) par. estimate stderror low95 % up95 % β0 86.41222 14.55656 53.28540 112.4174 β1 -14.89503 2.94585 -20.13681 -8.1682 σ 1.18943 0.25506 0.80843 1.8933 table 6. mean estimations for material#3. µ = 86.41222 − 14.89503 · log(loading) (12) figures 7, 8 and 9 illustrate the quantile analysis. the quantile is a measure of location in statistics. it represents a threshold. a certain amount of the value is below, the residual amount above this threshold. using the example of material#1 the threshold is 21675.91 cycles. that means here that for a loading of 132.5 there is a 50 % probability of failure. the blue dotted lines represent the 95 % confidence interval. jmp can also perform a custom estimate. for example, the failure probability at the endurance limit is calculated for each material. the results are given in figures 10, 11 and 12. in the case of material#3 the material can sustain 4.7*106 cycles at a loading of 93 and a failure probability of 5 %. at a loading of 93 and cycles of 406 the failure probability becomes 4.4 %. 6. conclusion due to the very large scatter of the measurement results, the determination of the most suitable model 4 vol. 3/2016 statistical evaluation of fatigue data of components figure 7. quantile analysis for material#1. figure 8. quantile analysis for material#2. figure 9. quantile analysis for material#3. 5 chi nghia chung, zoltan major acta polytechnica ctu proceedings figure 10. custom estimate for material#1. figure 11. custom estimate for material#2. figure 12. custom estimate for material#3. to fit the experimental data was a major challenge. in some cases, one gets very unsatisfactory graphical representations of the models. by the application of several fatigue models, at least one suitable model for each material could be found. by the implementation of statistical analysis also the large outliers could be included in the parameter estimation. jmp is a powerful tool for statistical analysis. especially, the analysis of the failure probability is a very important feature. the estimate in jmp within the experimental data range leads to slightly different prediction than that of the fatigue models. extrapolation with jmp out of range is highly dependent on the quality of the measurement data and therefore, does not always leads to plausible results. list of symbols a,b,c,d material parameter n number of cycles ∆σ loading amplitude ∆σ0 loading amplitude at endurance limit ∆σst loading amplitude at ultimate strength σ standard deviation β0 positional parameter in jmp β1 shape parameter in jmp µ estimation for the mean value references [1] t. k. anastasios p. vassilopoulos. fatigue of fiber-reinforced composites. springer verlag london limited, london, 2011. [2] m. v. dieter radaj. ermüdung grundlagen für ingenieure. springer verlag berlin, heidelberg, 1995, 2003, 2007. [3] a. f.-c. e. castillo. a unified statistical methodology for modeling fatigue damage. springer science + business media b. v., 2009. [4] u. heidelberg. http://www.physi.uni-heidelberg.de/einrichtungen/ ap/elearning/index.php/animationen/37-anpassungvon-funktionen-an-messdaten/52-die-methode-derkleinsten-fehlerquadrate 2013. universität heidelberg, 2013. [5] s. j. m. c. r. a. schneider. best practice guide on statistical analysis of fatigue data. twi cambridge, uk, 2003. [6] a. s. for testing, materials. statistical analysis of fatigue data. astm, 1981. [7] s. j. miller. the method of least squares. brown university, providence, ri 02912, us. [8] t. e. j. s. udo kuckartz, stefan rädiker. statistik eine verständliche einfürung. springer fachmedien wiesbaden gmbh, 2010. [9] a. f. hans diefenbacher. statistik eine verständliche einfürung. andreas frank & ventus publishing aps, 2006. [10] p. schmolck. methoden der reliabilitätschätzung. universität der bundeswehr münchen, 2007. [11] s. i. inc. jmp version 11 documentation. sas institute inc., 2013. 6 acta polytechnica ctu proceedings 3:1–6, 2016 1 introduction 2 fatigue models 2.1 the basquin model 2.2 the strohmeyer model 2.3 the weibull model 2.4 the bastenaire model 3 techniques of parameter fitting 3.1 the method of least squares (l2-norm) 3.2 the method of least absolute deviations (l1-norm) 4 reliability 5 evaluation of fatigue data 5.1 fatigue models 5.2 reliability analysis in jmp 6 conclusion list of symbols references 298 acta polytechnica ctu proceedings 1(1): 298–302, 2014 298 doi: 10.14311/app.2014.01.0298 the world space observatory -ultraviolet (wso-uv) space telescope; status update in 2013 ana i. gómez de castro1, boris shustov2, mikhail sachkov2 1aegora research group, universidad complutense de madrid, spain 2institute of astronomy of the russian academy of sciences, russia corresponding author: aig@ucm.es abstract this is a short primer and a brief update on the status of the world space observatory-ultraviolet (wso-uv) project dated in may 2013. wso-uv is a 170m primary space telescope equipped for ultraviolet imaging and spectroscopy that will be operational in 2017 hosting an open science program for the world-wide scientific community. keywords: astronomical instrumentation space astronomy ultraviolet astronomy. 1 introduction the ultraviolet (uv) range is fundamental for astrophysical investigations, since the resonance transitions of the most abundant species in the universe occur at these wavelengths/energies. the radiation cut-off at wavelengths shorter than 2800 å by the earth’s atmosphere makes uv astronomy only accessible from space. thus uv astronomy began with space exploration. after the copernicus mission, the international ultraviolet explorer (iue) was launched in 1978, becoming the first uv space observatory operated in real time; the iue allowed to carry out spectroscopic observations from 1150å to 3200å. later on, the far ultraviolet spectroscopic explorer (fuse) mission (19992007) opened the 900å -1000å spectral range for spectroscopic studies. the galaxy evolution explorer (galex, 2003-2011) has mapped, for the first time, the uv sky. as today, the hubble space telescope (hst) is the only operational mission in the uv range. hst is expected to last for a few more years. all these missions have amply demonstrated the feasibility and relevance of uv studies (see gómez de castro et al. 2006, gómez de castro & brosch 2009, for detailed compilations). the world space observatory-ultraviolet (wsouv) is an international space mission born as a response to the growing demand for uv facilities by the astronomical community. in the horizon of the next decade, the wso-uv will be the only two-meters class mission in the post-hst epoch which will guarantee access to uv wavelenghts. the project is managed by an international consortium led by the federal space agency roscosmos (russia). in this article, we briefly describe the wso-uv project, its general objectives and its main features. special emphasis is made on the ground segment and the instrument issis, the contributions of spain to the project. figure 1: the wso-uv space observatory. 2 the wso-uv scientific objectives the wso-uv is a multipurpose observatory on a geosynchronous orbit, which will provide data of large importance to investigate several open problems in astrophysics. the science drivers of the project are: • the study of the diffuse baryonic content in the universe and its chemical evolution – the main topics will be the investigation of baryonic content in warm and hot igm, of damped lyman-α 298 http://dx.doi.org/10.14311/app.2014.01.0298 the world space observatory -ultraviolet (wso-uv) space telescope; status update in 2013 systems, the role of starburts and the formation of galaxies. • the study of the formation and evolution of the milky way – the uv plays a particularly important role in the determination of energy inputs of the gas interacting with stars, and in the investigation of magnetic fields on star formation. the milky way history could be tracked through observations complementary to those obtained by the gaia mission. • the physics of accretion and outflows: the astronomical engines – this cathegory includes stars, black holes, interacting binaries, pre-main sequence stars an, in general, all those objects where accretion plays an important role in the evolution of the system. the efficiency and time scales of the phenomena will be studied, together with the role of the radiation pressure and the disk instabilities. • the investigation of the (extra)solar planetary atmospheres and astrochemistry in presence of strong uv radiation fields – the properties of the atmospheres of t tauri stars to study the environment where protoplanets grow. (see also gómez de castro et al. 2009) 3 the wso-uv mission and instrumentation the wso-uv telescope has an f/10 ritchey-chretien mounting with a primary diameter of 170 cm. wsouv has been thought as an observatory-type mission henceforth carrying instrumentation for uv imaging and spectroscopy (shustov et al. 2009, 2011). the wso-uv imaging and slitless spectroscopy instrument (issis) is a multipurpose instrument with a mode selector wheel that permits to carry out imaging and slitless spectroscopy in the 1150-3200 å spectral range. the instrument is equipped with two mcp detectors, with csi and cste photocathods for fuv and nuv observations, respectively. the wso-uv spectrographs (wuvs) consists of a set of three instruments: • the far uv high resolution spectrograph (vuves) that will permit to carry out echelle spectroscopy with resolution r∼50,000 in the 1150-1760 å range. it will be equipped with a photon-counting, micro channel plate (mcp) detector • the near uv high resolution spectrograph (uves) to carry out echelle spectroscopy with resolution r∼55,000 in the 1740-3050 å range. it will be equipped with a ccd detector to observe in the near uv. • the long slit spectrograph (lss) that will provide low resolution (r∼ 1000), long slit spectroscopy in the 11500-3050 å range. the spatial resolution will be 1 arcsec also, the width of the slit is 0.5 arcsec. the detector is a ccd cooled to −100o c to be sensitive to the far uv. figure 2: the instruments compartment in wsouv.the numbered sections correspond to: [1] optical bench with issis and wuvs mounted, [2] spectrographs, [3] cylinder inset of the instruments compartment, [4] protective cover of the instruments compartment and [5] heat pipes. (see also, sachkov 2010) prior to final tests, after the end of the construction phase, wso-uv instrumentation is expected to provide sensitivities similar to those of the hst instrumentation. the factor of 2 difference in the collecting surface between hst and wso-uv is compensated by the, much more efficient, high earth orbit of the wsouv, a geosynchronous orbit with inclination 51o. this will also allow to carry out efficiently monitoring programs. moreover, modern ”state of the art” ccd detectors will be used in the spectrographs. wso-uv expected launch date is 2016 and will be operational for five years with a possible extension to five years more. the space telescope is planned to be operated from two sites at madrid (ucm) and moscow (inasan) that will also host also the science and mission archives. the ground segment is being designed under a shared operations scheme. 299 ana i. gómez de castro, boris shustov, mikhail sachkov 3.1 issis design and expected performance the imaging and slitless spectroscopy instrument (issis) will be a key part of the wso-uv instrumentation. issis is the first uv imager to be flown to high earth orbit, above the earth geocorona. hence the uv background will be dominated by the zodiacal contribution and the diffuse galactic background due to dust-scattered starlight (murthy et al. 2010). the instrument has been designed to make full benefit of the heritage left by the galactic evolution explorer (galex) mission. galex has surveyed about 80% of the sky at uv wavelengths, providing for the first time a nearly complete view of the uv universe (martin et al. 2003, bianchi et al. 2011). however, galex spatial resolution was ∼ 4.2 arcsec and had very moderate spectroscopic capabilities. issis resolution will be ≤ 0.1 arcsec. the fine guiding system of the wsouv telescope will guarantee a high pointing stability (better than 0.1 arcsec at 3σ). moreover, issis will be equipped with gratings for slitless spectroscopy with resolution 500, in the full 1150-3200 spectral range. in imaging mode, issis effective area is about 10 times that of the galex imagers. figure 3: the layout of the imaging and slitless spectroscopy instrument (issis). the acronyms mark the location of: detectors (mcp), filter wheels (fw), pickup mirror mechanism (rm), calibration lamp (cl), mode selector mechanism (msm) and the mirrors m1 and m2. issis is designed to be an instrument for analysis of weak uv point sources or clumpy extended sources, especially those with well defined geometry. uv imaging instruments have been often equipped with prisms or very low dispersion grisms. the rapid decay of the resolution of prisms such as the available in the solar blind channel (sbc) of the advanced camera system (acs) makes very difficult its use to map extended line emission at wavelengths above some 1350 å. as the transmittance of narrow band filters in the far uv is ≤ 3% , integral-field low resolution spectroscopy is the main mean to map nebular emission. issis gratings will make feasible to use the powerful uv diagnostic tools to determine the location of dusty blobs and measure electron densities and temperatures. the instrument is located below the primary mirror and above the optical bench. this location imposes additional constraints to the design, in terms of weight and size: the maximum weight on the optical bench is 61.5 kg, and the full instrument has to be fit within a flat cylinder of height 17 cm. issis is fed by the central part of the beam but a pick-up mirror is required to fold the beam from the telescope adding one reflection. the final design is a compromise between the scientific requirements and the telescope/platform requirements. figure 3 shows the layout of the instrument. 4 wso-uv science programs the wso-uv will run three major science programs (see malkov et al. 2011 for details): • the core program includes the key scientific programs that will carry over the scientific objectives of the mission. the core program will be run for the first two years of the mission by the consortium building and operating wso-uv. • national programs: each country or funding body contributing to the project is entitled to receive a fraction of the observing time proportional to its contribution. after the third year of the project, 60% of the observing time will be awarded to these programs. national calls are expected to be issued for the national programs though they will be synchronized with the general project calls. guaranteed time for the instruments teams should be included in the national contributions. • open program to the world wide scientific community. this program will handle a 40% of the observing time after the 3rd year of the mission. targets of opportunity observations will be managed within these programs. 5 wso-uv ground segment (gs) the wso-uv gs is comprised of all the infrastructure and facilities involved in the preparation and execution of the wso-uv mission operations, which typically encompass real-time monitoring and control of the spacecraft, telescope and instruments as well as reception, processing and storage of the scientific data. in principle, there will be two complete gs systems: the russian 300 the world space observatory -ultraviolet (wso-uv) space telescope; status update in 2013 one will be located in moscow (lavochkin association and institute of astronomy of the ras), and the spanish one will be sited at madrid. the satellite operations will be shared between both ground control centers, transferring the mission control from one center to the other on a regular basis (lozano et al 2011). figure 4: basic layout of wso-uv ground segment. the science operations system and a fraction of the mission operations system are part of the spanish contribution to the wso-uv. the remote proposal system (rps), the science data processing system (sdps), the science archive (sa) and the scheduling systems are defined by the international science team composed by spanish and russian science support teams based at the universidad complutense de madrid (ucm) and russian science institute of astronomy of the russian academy of science (inasan). the science support team (st) is part of the man power of the gs, and is responsible of laying the foundation of and supervising all the operations related to the mission primary users: the scientists. at mission level, the st constitutes the core of the future wso-uv international observatory. a sumary of the high level definition documents, approved at mission level, for the development of the main science systems for wso-uv gs can be found in gómez de castro et al. 2011. 6 conclusions wso-uv is an international observatory that will grant access to the uv range in the post-hst era. at the time these proceedings are being written, the project is evolving into the construction phase. acknowledgement the spanish participation in the wso-uv project is being funded by the ministry of industry, energy and tourism (minetur). the scientific contribution to wso-uv is being funded by the ministry of economy and competitivity through the grant aya2011-29754c03. the authors wish to thank the large industrial and scientific team in russia and spain who are involved in the project for their dedication and expertise. references [1] bianchi, l., efremova, b., herald, j. et al. 2011, ap&ss, 335, 161 doi:10.1007/s10509-010-0581-x [2] gómez de castro, a.i., wamsteker, w., barstow, m. et al. 2006 ap&ss, 303, 133 doi:10.1007/s10509-006-9057-4 [3] gómez de castro, a.i., & brosch, n. 2009 ap&ss, 320, 1 doi:10.1007/s10509-009-0005-y [4] gomez de castro a.i. et al. 2009, in: chavez et al., new quests in stellar astrophysics. ii. ultraviolet properties of evolved stellar populations. p.319 springer, berlin [5] gómez de castro, a.i., sestito, p., sanchez doreste, n., et al. 2012, space ops 2012, on-line proceedings (http://www.spaceops2012.org/proceedings/ proceedings.html) [6] gómez de castro et al. 2012, space telescopes and instrumentation 2012: ultraviolet to gamma ray. proceedings of the spie, volume 8443, article id. 84432w. [7] lozano, j.m. et al. 2010, space ops 2010, aiaa 2010-2212 [8] malkov, o. et al., 2011 ap&ss, 335, 323 doi:10.1007/s10509-010-0589-2 [9] martin, c. et al. 2003, future euv/uv and visible space astrophysics missions and instrumentation. edited by j. chris blades, oswald h. w. siegmund. proceedings of the spie, volume 4854, pp. 336-350. [10] murthy, j., henry, r. c., & sujatha, n. v. 2010, apj, 724, 1389 doi:10.1088/0004-637x/724/2/1389 [11] reutlinger a. et al. 2011, apss 335, 311 [12] sachkov m. 2010, apss 329, 261 [13] shustov, b. et al. 2009 ap&ss, 320, 187 [14] shustov, b. et al. 2011 ap&ss, 335, 282 301 http://dx.doi.org/10.1007/s10509-010-0581-x http://dx.doi.org/10.1007/s10509-006-9057-4 http://dx.doi.org/10.1007/s10509-009-0005-y http://dx.doi.org/10.1007/s10509-010-0589-2 http://dx.doi.org/10.1088/0004-637x/724/2/1389 ana i. gómez de castro, boris shustov, mikhail sachkov discussion james beall: any idea of a possible launch date?. it’s a beautiful instrument. ana i. gomez de castro: the foreseen launch date is end of 2016. nino panagia: i believe this is a very important project. i’m wondering what is its expected lifetime. ana i. gomez de castro: the mission lifetime is 5 years plus 5 additional years after review. a 302 introduction the wso-uv scientific objectives the wso-uv mission and instrumentation issis design and expected performance wso-uv science programs wso-uv ground segment (gs) conclusions 170 acta polytechnica ctu proceedings 2(1): 170–173, 2015 170 doi: 10.14311/app.2015.02.0170 new insights from inside-out doppler tomography e. j. kotze1,2, s. b. potter1 1south african astronomical observatory, po box 9, observatory 7935, cape town, south africa 2astrophysics, cosmology and gravity centre (acgc), department of astronomy, university of cape town, private bag x3, rondebosch 7701, south africa corresponding author: ejk@saao.ac.za abstract we present preliminary results from our investigation into using an “inside-out” velocity space for creating a doppler tomogram. the aim is to transpose the inverted appearance of the cartesian velocity space used in normal doppler tomography. in a comparison between normal and inside-out doppler tomograms of cataclysmic variables, we show that the inside-out velocity space has the potential to produce new insights into the accretion dynamics in these systems. keywords: accretion accretion discs methods: spectroscopic binaries: close dwarf novae polars cataclysmic variables. 1 introduction cataclysmic variables (cvs) are quintessential stellar objects for the study of mass transfer, accretion flows and accretion discs. the typical interacting system consists of a secondary low-mass, red dwarf star which is filling its roche lobe and is transferring mass via the inner lagrangian point (l1) onto the primary white dwarf star (see warner 1995 for a comprehensive review). doppler tomography, as introduced by marsh & horne (1988), is aimed at rendering the information locked-up in phased-resolved spectra of a cv into a twodimensional map of the binary components in velocity space (doppler tomogram). 2 doppler tomography 2.1 spatial coordinates fig. 1 shows a model cv with an accretion disc in a cartesian spatial coordinate frame that co-rotates with the system. as described by marsh & horne (1988), this two-dimesional frame has its origin at the centre of mass of the system [marked with a plus (+)], the x-axis along the line connecting the centres of mass of the primary and secondary [marked with crosses (×)], and the y-axis parallel to the velocity vector of the secondary. the orbital motion is assumed to be counter-clockwise around the centre of mass of the system. the model cv assumes the following system parameters: inclination i = 87◦; mass of the primary m1 = 0.8; mass ratio q = m2/m1 = 0.2 and orbital period porb = 0.083333 days (120 minutes). the inner disc radius is derived using these parameters and assuming a keplerian velocity of ∼ 2.37 × 103 km s−1. the 3:1 resonance radius is taken to be the outer disc radius. secondary accretion stream inner edge of disc outer edge of disc roche lobe of primary a(0.02,-0.12) b(-0.62,0.0) x y figure 1: co-rotating cartesian spatial coordinate frame for a model cv. 2.2 velocity coordinates the left and middle panels of fig. 2 show, respectively, the two-dimensional cartesian and polar velocity coordinate frames which correspond to the co-rotating spatial frame shown in fig. 1, with overlays of the velocity profiles of all the main components of the model cv. the polar velocity coordinate frame is obtained by either transforming the cartesian spatial frame to a polar spatial frame which is then projected into a polar velocity frame or by transforming the cartesian velocity frame into a polar velocity frame. since we found that a polar frame allows for easier transformation of the circularly symmetric velocity profile of a doppler tomogram, this is the first step in establishing an inside-out layout. 170 http://dx.doi.org/10.14311/app.2015.02.0170 new insights from inside-out doppler tomography v(r′,θ) [103 km s-1, degrees] secondary accretion stream inner edge of disc outer edge of disc roche lobe of primary a(0.86,55.2°) b(0.73,270°) 0.0 1.0 2.0 r′ θ v(r′,θ) v(r,θ) [103 km s-1, degrees] secondary accretion stream inner edge of disc outer edge of disc roche lobe of primary a(0.86,55.2°) b(0.73,270°) 1.0 2.0r θ v(r,θ) -2.0 -1.0 0.0 1.0 2.0 -2.0 -1.0 0.0 1.0 2.0 v y [ 1 0 3 k m s -1 ] v x [10 3 km s -1 ] secondary accretion stream inner edge of disc outer edge of disc roche lobe of primary a(0.49,0.70) b(0.0,-0.73) figure 2: cartesian (left), polar (middle) and inside-out polar (right) velocity space. 3 inside-out doppler tomography the right panel of fig. 2 shows the same model cv presented in the left and middle panels, but with the zero velocity origin and the maximum velocities transposed, creating an inside-out polar velocity space. the most notable aspects of the inside-out frame are the inner and outer edges of the accretion disc appearing the “right” way around while the secondary and ballistic stream are outside the disc. the secondary appears upside down because it is orbiting as a solid body, i.e., the outside is moving faster than the inside. the ballistic stream also “curves” inwards as it accelerates towards the disc and primary. 4 examples the example doppler tomograms have been created in polar and inside-out polar velocity space using a modified version of the fast maximum entropy doppler mapping code presented by spruit (1998). 4.1 spiral shocks in the accretion disc of the dwarf nova ip peg fig. 3 shows comparative non-axisymmetric normal and inside-out tomograms for the heii 4686å emission line from phase-resolved spectroscopy of ip peg. the model velocity overlays were calculated using porb = 0.158206 days (∼ 228 minutes), q = 0.48, m1 = 1.16 and i = 83.8◦ (copperwheat et al. 2010). the inner and outer edges of the disc (solid lines), the roche lobes of the primary (dashed line) and secondary (solid line) as well as a single particle ballistic trajectory (solid line) from l1 to 20 ◦ in azimuth around the primary, are shown. in the normal tomogram the secondary appears as a bright spot inside the disc, whereas it becomes a diffuse patch (spread over more pixels) outside the disc in the inside-out tomogram. however, in the inside-out tomogram the disc appears the “right” way around as the two spiral shocks appear to be spiralling “inward” towards higher velocities. 4.2 ballistic and magnetic accretion flow in the polar hu aqr comparative normal and inside-out tomograms for the heii 4686å emission line from phase-resolved spectroscopy of hu aqr are shown in fig. 4. porb = 0.086820 days (∼ 125 minutes), q = 0.4, m1 = 0.875 and i = 84◦ (one of the models from schwope et al. 1997) were used to calculate the model velocity overlays. the roche lobes of the primary (dashed line) and secondary (solid line) as well as a single particle ballistic trajectory (solid line) from l1 to 65 ◦ in azimuth around the primary, are shown. a dipolar axis azimuth and co-latitude of ∼ 38◦ and ∼ 12◦ (heerlein et al. 1999) respectively, were used to calculate magnetic dipole trajectories (thin dotted lines) at 10◦ intervals from 15◦ to 65◦ in azimuth around the primary. the secondary appears as a bright spot in both the normal and the inside-out tomograms. in the normal tomogram the ballistic part of the accretion flow appears as a prominent ridge with an apparent consistent brightness from l1 to 1.0×103 km s−1, but with almost no discernible detail at higher velocities. in the insideout tomogram the ballistic flow is more exposed and varying in brightness, but retaining discernible detail to at least 1.5 × 103 km s−1. low-velocity (0.0 − 0.5 × 103 km s−1) emission associated with the magnetic coupling region is seen as a diffuse patch in the lower left quadrant of both tomograms. there is no high-velocity (> 1.5 × 103 km s−1) emission discernible in the lower left quadrant of the normal tomogram, whereas in the inside-out tomogram there is a small patch of emission between 1.5 − 2.0 × 103 km s−1 that can be linked to the magnetically confined accretion flow close to the primary. 171 e. j. kotze, s. b. potter 5 conclusions in a normal tomogram the lower-velocity features tend to dominate the brightness scale since they are concentrated over fewer pixels compared to higher-velocity features. in an inside-out tomogram the converse is true with teneous higher-velocity features being enhanced while prominent lower-velocity features are more spread out and exposed. teneous lower-velocity features, however, may be diluted to the point of becoming indiscernible similar to teneous higher-velocity features in a normal tomogram. we have shown that inside-out doppler tomography projects the accretion disc of a cv the “right” way around with the ballistic stream and the secondary outside the disc. furthermore, the accretion flow of a polar appears more intuitive in an inside-out tomogram with the ballistic stream curving “inwards” and the magnetic flows being more exposed. therefore, we conclude that our new technique of inside-out doppler tomography is complementary to the existing technique. acknowledgement the authors would like to thank axel schwope and danny steeghs for sharing their data of hu aqr and ip peg, respectively. references [1] copperwheat c.m., et al.: 2010, mnras, 402, 1824. doi:10.1111/j.1365-2966.2009.16010.x [2] heerlein c., horne k., schwope a.d.: 1999, mnras, 304, 145. doi:10.1046/j.1365-8711.1999.02311.x [3] marsh, t.r., horne, k.: 1988, mnras, 235, 269. doi:10.1093/mnras/235.1.269 [4] schwope a.d., mantel k.-h., horne k.: 1997, a&a, 319, 894. [5] spruit h.c.: 1998, arxiv:astro-ph/9806141. [6] warner b.: 1995, cambridge astrophysics series 28, cataclysmic variable stars. cambridge univ. press, cambridge figure 3: doppler tomography of ip peg. normal (top) and inside-out (bottom) tomograms are shown for comparison. the input and reconstructed trailed spectra are shown in the left and right panels, respectively. 172 http://dx.doi.org/10.1111/j.1365-2966.2009.16010.x http://dx.doi.org/10.1046/j.1365-8711.1999.02311.x http://dx.doi.org/10.1093/mnras/235.1.269 new insights from inside-out doppler tomography figure 4: doppler tomography of hu aqr. normal (top) and inside-out (bottom) tomograms are shown for comparison. the input and reconstructed trailed spectra are shown in the left and right panels, respectively. discussion dmitry kononov: how does your new technique redistributes brightness and are there some physical suppositions behind this redistribution? enrico kotze: as we see in the example of ip peg the upper bright spiral shock in the normal tomogram appears less bright in the inside-out tomogram since it is projected over a larger area of the image. similar to the normal technique the brightness distribution in the new technique is purely a function of the projection of the spectra into the velocity space frame and we have no claims that this is a representation of the physical brightness distribution of the system. linda schmidtobreick: do you encounter problems with the resolution as the low velocities are now spread over a large circle? enrico kotze: yes, there can be a loss of resolution at low velocities. this is the reverse of what happens in the normal tomograms where high velocities are spread over a larger area and we can encounter a loss of resolution in the high-velocity features. that is why we feel the inside-out technique is complementary to the normal technique. where the normal technique tends to enhance low-velocity features with some loss in the resolution of high-velocity features, the inside-out technique tends to enhance high-velocity features with some loss in the resolution of low-velocity features. david buckley: is the magnetic longitude of the accreting pole(s) a parameter in the inside-out maps? enrico kotze: yes, for the model velocity profile overlays in the inside-out tomograms of polars we take both the azimuthal and longitudinal inclination of the assumed magnetic dipole into account. 173 introduction doppler tomography spatial coordinates velocity coordinates inside-out doppler tomography examples spiral shocks in the accretion disc of the dwarf nova ip peg ballistic and magnetic accretion flow in the polar hu aqr conclusions 293 acta polytechnica ctu proceedings 2(1): 293–296, 2015 293 doi: 10.14311/app.2015.02.0293 non relational models for the management of large amount of astronomical data b. l. martino1, m. federici 2 1associated inaf iaps 2istituto di astrofisica e planetologia spaziali, inaf iaps via fosso del cavaliere 100, 00133 roma, italy corresponding author: brunolmartino@iasi.cnr.it abstract the objective of this paper is the comparison between two different database typologies: the relational and the nonrelational architecture, in the context of the applications related to the use and distribution of astronomical data. the specific context is focused to problems quite different from those related to administrative and managerial environments within which were developed the leading technologies on which are based the modern systems of massive storage of data. the data provided by astronomical instrumentation are usually filtered out by the front-end system (trigger, anticoincidence, dsp etc.), so they do not require special controls of congruence. moreover, the related storage systems must be able to ensure an easy growth, minimizing human systemistic interventions and automating the related actions. the use of a non-relational architecture (nosql), offers great advantages during the insertion of informations within a data base, while the response speed of the queries is mainly tied to their type and complexity. keywords: dbms mongodb nosql mysql gsc catalog. 1 using dbms a careful planning of the use of a file system allows to store informations in a rational way but, whatever the criterion used to organize an archive based only on files, it is not possible to build a search system that can guarantee sufficient flexibility. the user of a database, conceived in this manner, is required to know in detail its structure and it’s organization. almost any application focused on the analysis of astronomical data may read data in fits format (wells et al., 1981) fits format allows to add to a collection of data a set of additional information used to allow their better characterization. a dbms (data base management system) is a software infrastructure designed to operate on large data sets with the goal of optimizing the data: • storage • access • sharing • protection a dbms allows, through the use of its command language, the imposition of constraints of consistency, the creation of indexes to improve performance and retrieval of data independently from their physical representation. well-known examples of the use of advanced databases in high-energy physics are represented by spires (bourne et al., 2003), opera (agafonova et al., 2009) and !chaos (bisegni, 2012) and is consolidated practice their adoption as a storage medium in the ground segment. 1.1 relational models the relational databases today are the most popular; the model they are based it was proposed by codd (1970). the relational model allows access to data at the logical level by providing a complete independence from their physical organization. the logical representation of the data in the relational model is based on the concept of relationship, in algebraic terms; it is common to use the term ”table” in place of ”algebraic relationship” and the term ”relationship” to indicate an association between the data. the language sql (date et al., 1997) is the standard language ”de facto” for defining, manipulating and interrogation of relational databases; this is a declarative language and not procedural. in the relational model a logical unit of work is defined transaction and is constituted by a sequence of operations of reading and writing, which must meet some properties, known as acid property (from atomicity, consistency, isolation, durability). atomicity: a transaction is an atomic unit of operations, validated or canceled depending on whether or not they reach a successful conclusion (rollback / commit); consistency: at 293 http://dx.doi.org/10.14311/app.2015.02.0293 b. l. martino, m. federici the end of the transaction if the initial state is correct, even the final state must be; isolation: the action of a transaction should not interfere with each other; duration: the effects of a transaction must be persistent; the most popular relational databases are oracle (kunh, 2010) and mysql (schwartz, 2012)). 1.2 non relational models recently, have been developed a series of new dbms systems, to provide an high horizontal scalability, in order to achieve high performance in the read / write operations of database distributed across multiple servers geographically delocalized (cloud). many of these new systems are called nosql data stores. the definition of nosql, which stands for ”not only sql” or ”not relational”, was used for the first time in 1998 for an open source relational database. the non-relational systems, do not attempt to provide the classic acid guarantees, typical of relational databases, but embrace the model b.a.s.e. (basic available, soft-state, eventual consistency) where some constraints are relaxed: basic available: the database can operate even if a part is no longer available; soft-state: the system status may change over time even in absence of input data; eventual consistency: the data may not be updated immediately, but will be consistent throughout the system within a finite time as stated by the cap theorem (lynch et al., 2002) can not be achieved consistency, availability and partition tolerance at the same time but only two of these features at a time. the systems based on nonrelational architecture follow the base model and are those that allow to overcome the major limitation of the rdbms (relational data base management system): the scalability. in many areas, consistency and/or availability offered by relational databases are not essential, e.g. in astronomy and astrophysics applications. in a highly available and tolerant partitioning system, alteration of data base will reach all nodes not instantly but within a finite time, if a reading is done on a node is not synchronized with the last write, it returns the last valid data (stale data). 2 dbms nosql a property of nosql database is to be free of pattern (schemaless), with consequent advantages and disadvantages like ease of deployment, but sometimes, more difficulties to construct complex queries. the nonrelational logical models not have the same expressive power of the relational model and can be classified into four main families: key-value, column-family, document store and graph. the most common are the mongodb nosql architecture (chodorow, 2011), hbase (george, 2011) and cassandra (lakshman et al., 2010); their characteristics are summarized in figure 1. mongodb is a document-oriented database, which is based on aggregates, which may have a structure with multiple hierarchical levels and groups that can be variously indexed. mongodb uses the javascript language, which is inherently single-threaded while cassandra uses the language cql (cassandra query language, simplified variant of sql), which allows the management of distributed databases. both rely on the file system of the machine on which they were installed without introducing data abstractions. hbase is a distributed database (based on the project hadoop (sammer, 2012) written in java, that uses storage devices located on different hosts also geographically distributed, interconnected via networks. hbase uses hdfs, a portable and scalable distributed file system initially developed for the framework hadoop. figure 1: key features of nosql architectures. to carry out complex operations is used the mapreduce paradigm (dean et al., 2004), which allows to divide the computation in many elaborations of lower complexity to achieve the processing of large datasets in parallel on multiple cores/cpu/computer. 3 mysql vs mongodb in order to evaluate the possible use with advantage of non-relational architectures in astronomy and astrophysics area we realized the mdbirs system (showed in figure 2), composed by 10 pc, equipped with intel i7, 16gbytes of ram, 2 hd 1tbytes connected via ethernet lan at 1gbit of speed figure 2: the mdbirs system. 294 non relational models for the management of large amount of astronomical data in order to assess objectively the behavior of the system tests were carried out so as not to make use of the advanced characteristics of the query languages used. for this reason it was chosen as the working set of data the gsc catalog (stars general catalog). gsc is composed of a single table and is used primarily to provide support for the planning and guiding stars of the hst observations (dalcanton, 2009), the jwst (gardner at al., 2006), gaia (busso et al., 2012) and some groundbased telescopes of new generation. furthermore, the services provided by the machines on which they are installed databases are not affected by processes (user or system) not strictly essential to their functioning and the software installed on the machines which host the test has been aligned to the same versions. the databases under test are the two most widely used open source architectures in the world, mysql for its versatility, speed and diffusion and mongodb for its attitude to the horizontal growth (scalability) and its robustness. starting from the data available, have been generated some of the series of samples of increasing amplitude, so as to highlight the behavior of the two systems both with regard to populate the database, than to retrieve the information of interest. the test was performed using queries significant from the point of view of the community of users, which allow to highlight the behavior of databases in real operating conditions. in particular: • query1: selection on healpix (mapping system applied to the celestial sphere) • query 2: selection of objects in a spatial region which satisfy some conditions in magnitude and color • query 3: selection of non-stellar objects which meet a condition in color • query 4: computing of the average magnitudo, resolution of one degree square 4 test results the database engine installed by default in mysql since version 5.0 (innodb) is transactional, acid compliant and uses the row-level locking strategy (constraint of exclusive use for the time necessary to perform the required actions). the data entry test was made from data provided in csv (text data whose fields are marked by the separator character comma.) as shown by the graph in figure 3, the behavior of mongodb is faster than mysql, thanks to the absence of checks carried out on the integrity of the data; information are stored in files as key-value pairs. figure 3: insert performances vs data set size. regarding the execution time of the test queries, the results are shown in the graph of figure 4 and are substantially comparable. in the case of the gsc catalog , the occupation of the relevant mysql table is about 116 gbytes of data collection while the corresponding mongodb occupies about 657 gbytes (the size is 5 times bigger). the tables were not associated with indexes, so searches are performed on the whole set of data with a time dependent to amplitude of the set itself. figure 4: queries performances vs data set size. 5 future activities mongodb is a good choice to obtain good safety and performances using local data sharding. our goal is to achieve a full geographic data delocalization (cloud) in order to obtain: continuity of access to data (total disaster recovery) to the user community optimized average access time we evaluated two possible scenarios: hbase: hadoop based nosql data base cassandra: the apache cassandra database hbase is a masterslave system including two types of machines: hmaster: access control hregionserver: local data replication by using the stargate plugin can be achieved local caches able to speed up response times by minimizing the network traffic to the master. every region server keeps a copy of the data so the system ensures very 295 b. l. martino, m. federici high reliability. it is a nosql solution initially developed by facebook which has in the p2p architecture its focal point. 6 conclusions mongodb has proven much more efficient and fast in data entry and is particularly suited to the management of flows of data to be stored without downtime. in the configuration we used the rate of loading is about 15000 rows/sec (the average rows length is 256 bytes). by using more powerful hardware such as high speed network interfaces (10 gbit/sec or higher) it’s possible to greatly increase the speed of data acquisition. regarding the search of data in the database the results show a behavior strongly dependent on the number of records and the presence or absence of critical elements such as complex calculations. the command language of mongodb does not contain any advanced mathematical primitives as in the case of mysql, which can rely on a large library of mathematical functions. the use of mongodb on structures designed to be handled by a rdbms is inappropriate because it does not exploits the potential of its schemaless organization (nosql). acknowledgement we thanks paola marrese by asdc-asi for providing us the test data, , the director of iaps-inaf prof pietro ubertini, and a special thanks to franco giovannelli and his staff for allowing us to participate to this workshop references [1] wells, d. c., greisen, e. w., and harten, r. h.: 1981, a flexible image transport system. astronomy & astrophysics supplement series, 44, 363-370. [2] bourne, c. p., hahn, t. b.: 2003, a history of online information services, 1963-1976, mit press. [3] agafonova, n., et al.: 2009, the detection of neutrino interactions in the emulsion/lead target of the opera experiment, journal of instrumentation. doi:10.1088/1748-0221/4/06/p06020 [4] bisegni, c.: 2012, thesis in computer science, university tor vergata, rome. [5] codd, e.f.: 1970, a relational model of data for large shared data banks, communications of the acm 13 (377-387). doi:10.1145/362384.362685 [6] c. j. date, c. j., darwen, h.: 1997, a guide to the sql standard: a users guide to the standard database language sql, addison wesley. [7] kuhn, d.: 2010, pro oracle database 11g administration, apress. doi:10.1007/978-1-4302-2971-1 [8] schwartz, b.: 2012, high performance mysql: optimization, backups, and replication, o’reilly. [9] lynch, n., gilbert, s.: 2002, brewer’s conjecture and the feasibility of consistent, available, partition-tolerant web services, acm sigact news, volume 33 issue 2, pg. 51-59. [10] chodorow, k.: 2011, scaling mongodb, o’reilly. [11] george, l: 2011, hbase: the definitive guide, o’reilly. [12] lakshman, a., malik, p.: 2010, cassandra: a decentralized structured storage system, acm sigops operating systems review, vol. 44, no.2. doi:10.1145/1773912.1773922 [13] sammer, e.: 2012, hadoop operations, o’reilly. [14] dean, j., ghemawat, s.: 2004, mapreduce: simplified data processing on large clusters, in osdi’04: sixth symposium on operating system design and implementation, san francisco, ca, december. [15] dalcanton, j.j.: 2009, 18 years of science with the hubble space telescope, nature jan 1;457(7225):41-50. [16] gardner, j. p., at al.: 2006, the james webb space telescope, space science reviews. doi:10.1007/s11214-006-8315-7 [17] busso, g., de angeli, f., montegriffo, p.: 2012, the gaia photometric data processing, in proc. spie 8442, space telescopes and instrumentation 296 http://dx.doi.org/10.1088/1748-0221/4/06/p06020 http://dx.doi.org/10.1145/362384.362685 http://dx.doi.org/10.1007/978-1-4302-2971-1 http://dx.doi.org/10.1145/1773912.1773922 http://dx.doi.org/10.1007/s11214-006-8315-7 using dbms relational models non relational models dbms nosql mysql vs mongodb test results future activities conclusions 274 acta polytechnica ctu proceedings 1(1): 274–277, 2014 274 doi: 10.14311/app.2014.01.0274 multiscale modeling of astrophysical jets james h. beall1,2,3, john guillory* 3, david v. rose4, michael t. wolff2 1st. johns college, annapolis, md 2space sciences division, naval research laboratory, washington, dc; 3college of sciences, george mason university, fairfax, va; 4voss scientific, albuquerque, nm; corresponding author: beall@sjc.edu abstract we are developing the capability for a multi-scale code to model the energy deposition rate and momentum transfer rate of an astrophysical jet which generates strong plasma turbulence in its interaction with the ambient medium through which it propagates. we start with a highly parallelized version of the vh-1 hydrodynamics code (coella and wood 1984, and saxton, et al., 2005). we are also considering the pluto code (mignone et al. 2007) to model the jet in the magnetohydrodynamic (mhd) and relativistic, magnetohydrodynamic (rmhd) regimes. particle-in-cell approaches are also being used to benchmark a wave-population models of the two-stream instability and associated plasma processes in order to determine energy deposition and momentum transfer rates for these modes of jet-ambient medium interactions. we show some elements of the modeling of these jets in this paper, including energy loss and heating via plasma processes, and large scale hydrodynamic and relativistic hydrodynamic simulations. a preliminary simulation of a jet from the galactic center region is used to lend credence to the jet as the source of the so-called the fermi bubble (see, e.g., su, m., & finkbeiner, d. p., 2012) *it is with great sorrow that we acknowledge the loss of our colleague and friend of more than thirty years, dr. john ural guillory, to his battle with cancer. keywords: jets active galaxies blazars intracluster medium non-linear dynamics plasma astrophysics. 1 introduction recent high-resolution (vlba) observations of astrophysical jets (see, e.g., lister et al. 2009) reveal complex structures apparently caused by ejecta from the central engine as they interacts with both surrounding interstellar material such as broad-line region (blr) and narrow-line region (nrl) clouds, and ejecta from prior episodes of activity. a particularly trenchant example of these complex interactions is also shown by the galactic microquasar, sco x-1 (fomalhaut, geldzahler, and bradshaw, 2001). such observations can be used to inform models of the jet-ambient-medium interactions. based on an analysis of these data, we posit that a significant part of the observed phenomena come from the interaction of the ejecta with prior ejecta as well as interstellar material. 2 scales of jet interactions with the ambient medium large scale hydrodynamic simulations of the interaction of astrophysical jets with the ambient medium through which they propagate can be used to illuminate a number of interesting consequences of the jets’ presence. these include acceleration and entrainment of the ambient medium, the effects of shock structures on star formation rates, and other effects originating from ram pressure and turbulence generated by the jet (see, e.g., basson and alexander, 2002; zanni et al. 2005; and krause and camenzind 2003; perucho, et al. 2012). we have presented results for large scale hydrodynamic simulations and initial relativistic hydrodynamic simulations in previous works (beall et al., 1999, 2003, 2006). as noted in those papers, magneto-hydrodynamic (mhd) and (rmhd) simulations neglect important species of physics: the microscopic interactions that occur because of the effects of particle-particle interactions and the interactions of particles with the collective effects that accompany a fully or partially ionized ambient medium (i.e. a plasma). while the physical processes (including plasma processes) in the ambient medium can be modeled in small regions by pic (particle-in-cell) codes for some parameter ranges, simulations of the larger astrophysical jet structure with such pic codes are not possible with current or foresee274 http://dx.doi.org/10.14311/app.2014.01.0274 multiscale modeling of astrophysical jets able computer systems. for this reason, we have modeled these plasma processes in the astrophysical regime by means of a system of coupled differential equations which give the wave populations generated by the interaction of the astrophysical jet with the ambient medium through which it propagates. a detailed discussion of these efforts can be found, variously, in scott et al. (1980), rose et al. (1984), rose et al. (1987), beall (1990), and beall et al. (2003). the scales of these interactions range from kilometers to kiloparsecs. 2.1 energy loss, energy deposition rate, and momentum transfer from plasma processes the system of equations used to determine the normalized wave energy densities is very stiff. scott et al. (1980) estimated the equilibrium solution of this system of equations for heating of clusters of galaxies, and rose et al. (1984) and beall (1990) showed dynamical solutions that confirmed the stability of the equilibrium solutions. solving the system of equations yields a time-dependent set of normalized wave energies (i.e., the ratio of the wave energy divided by the thermal energy of the plasma) that are generated as a result of jets interaction with the ambient medium. these solutions can yield an energy deposition rate (de/dt), an energy deposition length (de/dx), and ultimately, a momentum transfer rate (dp/dt (1/vb) ∗ (de/dt) that can be used to estimate the effects of plasma processes on the hydrodynamic evolution of the jet. for this part of the analysis, we suppose that a portion of the jet is composed relativistic particles of either e±, p − e−, or more generally, a charge-neutral, hadron-e− jet, with a significantly lower density than the ambient medium. the primary energy loss mechanism for the electron-positron jet is via plasma processes, as beall (1990) notes. kundt (1987, 1999) also discusses the propagation of electron-positron jets. beall et al., (2006) illustrate two possible solutions for the system of coupled differential equations that model this mode of the jet-ambient medium interaction: a damped oscillatory and an oscillatory solution. the landau damping rate for the two-temperature thermal distribution of the ambient medium is used for these solutions. the average energy deposition rate, < d(α�1)/dt >, of the jet energy into the ambient medium via plasma processes can be calculated as < d(α�1)/dt >= npkt < w > (γ1/ωp)ωpergs cm −3s−1, where k is boltzmann’s constant, t is the plasma temperature, < w > is the average (or equilibrium) normalized wave energy density obtained from the wave population code, γ1 is the initial growth rate of the two-stream instability, and α is a factor that corrects for the simultaneous transfer of resonant wave energy into nonresonant and ion-acoustic waves. the energy loss scale length, deplasma/dx = −(1/nbvb)(dα�1/dt), can be obtained by determining the change in γ of a factor of 2 with the integration ∫ dγ = − ∫ [d(α�1)/dt]dl/(vbnbm ′ c2) as shown in rose et al., 1978 and beall 1990, where m ′ is the mass of the beam particle. thus, lp = ((1/2)γcnbmc 2)/(dα�1/dt) cm is the characteristic propagation length for collisionless losses for an electron or electron-positron jet, where dα�1/dt is the normalized energy deposition rate (in units of thermal energy) from the plasma waves into the ambient plasma. in many astrophysical cases, this is the dominant energy loss mechanism. we can therefore model the energy deposition rate (de/dt) and the energy loss per unit length (de/dx), and ultimately the momentum loss per unit length (dp/dx) due to plasma processes. figure 1: simulation using a highly-parallelized version of the vh-1 hydrodynamics code. the figure shows the transverse cross-section (figure 1a) and the detail of the jet head in the y-z cross-section (figure 1b) for a fully 3-dimensional hydrodynamic simulation of jet with v = 1.5x109 cm/sec. the cross section at z = 300 for a jet with v = 1.0x1010 cm/s. the cross section represents approximately 2 kiloparsecs from the jet origin. we have used the 5123 simulation in this representation for ease of presentation, given the size and difficulty of displaying the 20483 data. note the well-developed rayleigh-taylor instabilities at the jet-ambient medium boundary in figure 1a. beall, guillory, and rose (2009) have compared the results of a pic code simulation of an electronpositron jet propagating through an ambient medium of an electron-proton plasma with the solutions obtained by the wave population model code, and have found good agreement between the two results (see figure 1 from that paper). at the same time, that paper demonstrates that the ambient medium is heated and entrained into the jet. that analysis also shows that a relativistic, low-density jet can interpenetrate an ambient gas or plasma. initially, and for a significant fraction of its propagation length, the principal energy loss mechanisms for such a jet interacting with the ambient medium is via plasma processes (rose et al. 1984, beall 1990). 275 james h. beall et al. as part of our research into the micro-physics of the interaction of jets with an ambient medium, we continue to investigate the transfer of momentum from the jet, and expect to present these results shortly. in order to proceed to a more detailed analysis of the issue of momentum transfer, we have used modern pic code simulations to study the dynamics of caviton formation, and have confirmed the work of robinson and newman (1990) in terms of the cavitons’ formation, evolution, and collapse. 2.2 results of hydrodynamical calculations figure 1 shows two views of a simulation for an astrophysical jet for parameters appropriate to a seyfert galaxy, using a highly-parallelized version of the vh1 hydrodynamics code. the figure shows the transverse cross-section (figure 1a) and the detail of the jet head in the y-z cross-section (figure 1b) for a fully 3-dimensional hydrodynamic simulation of jet with v = 1.5x109 cm/sec. the cross section at z = 300 for a jet with v = 1.0x1010 cm/s. the cross section occurs at approximately 2 kiloparsecs from the jet origin. figure 2: simulation using the vh-1 code of a supersonic jet with parameters appropriate to the galactic center region. in this simulation, also, we use the vh-1 code for a supersonic jet in the x-y pressure cross section (1a) and with an x-y density cross section (1b) for a supersonic jet with v = 2.0x108 cm/s. such a jet originating from the supermassive black hole at the center of the milky way could account for the production of the gamma-ray structure known as the galactic center bubble found by the fermi satellite. the vertical scale of the jet represents approximately 2 kiloparsecs from the jet origin. we have used the 5123 simulation in this representation for ease of presentation, given the size and difficulty of displaying the 20483 data figure 2 shows a simulation of a supersonic jet for parameters appropriate to the galactic centre region. in this simulation, also, we use the vh-1 code for a supersonic jet in the x-y pressure cross section (1a) and with an x-y density cross section (1b) for a supersonic jet with v = 2.0x108 cm/s. such a jet originating from the supermassive black hole at the center of the milky way could account for the production of the gamma-ray structure known as the galactic center bubble found by the fermi satellite. the vertical scale of the jet in figure 2a represents approximately 2 kiloparsecs from the jet origin. we have used the 5123 simulation in this representation for ease of presentation. the detail of the jet-head structure (figure 2b) shows complex, transverse shock structures at the leading edge of the jet. the simulations we have run using 20483 cells confirm these features. 3 concluding remarks the effects of collective and particle processes, including plasma effects, can have observational consequences. beall (1990) has noted that plasma processes can slow the jets rapidly, and beall and bednarek (1999) have shown that these effects can truncate the low-energy portion of the γ-rays spectrum (see their figure 3), a similar effect will occur for particle-particle productions of neutrinos, pions, and (perhaps) neutrons. this could also reduce the expected neutrino flux from agn. the presence of plasma processes in jets can also greatly enhance line radiation species by generating high-energy tails on the maxwell-boltzmann distribution of the ambient medium, thus abrogating the assumption of thermal equilibrium. an analytical calculation of the boost in energy of the electrons in the ambient medium to produce such a high energy tail, with ehet ∼ 30 − 100kt, is confirmed by pic-code simulations. aside from altering the landau damping rate, such a high-energy tail can greatly enhance line radiation over that expected for a thermal equilibrium calculation (see beall et al. 2006, and beall, guillory, and rose (1999) for a detailed discussion). we are in the process of setting up runs for the pluto code (mignone et al. 2007) in order to benchmark its results with our results of the vh-1 code in appropriate parameter ranges. acknowledgement jhb and mtw gratefully acknowledge the support of the office of naval research for this project. thanks also to colleagues at various institutions for their continued interest and collaboration, including kinwah wu, curtis saxton, s. schindler and w. kapferer, and s. colafrancesco. references [1] basson, j. f., and alexander, p., 2002, the longterm effect of radio sources on the intracluster medium, mnras, 339, 353. 276 multiscale modeling of astrophysical jets [2] beall, j. h. et al., 1978, radio and x-ray variability of the nucleus of centaurus a (ngc 5128), ap.j., 219, 836. doi:10.1086/155845 [3] beall, j. h. and rose, w. k., 1981, on the physical environment in 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[15] lister, m. l. et al., 2009, astronom. j., 137, 3718. doi:10.1088/0004-6256/137/3/3718 [16] mignone, a. et al., 2007, ap.j. suppl., 170, 228. [17] newman, d. l., winglee, r. m., robinson, p. a., glanz, j., and goldman, m. v., 1990, simulation of the collapse and dissipation of langmuir wave packets, phys. fluids b, 2, 2600. doi:10.1063/1.859385 [18] perucho, m., 2012, jets in high-mass microquasars, mem. s. a. it., 83, 297. [19] rose, w. k., guillory, j., beall, j. h., and kainer, s., 1984, the interaction of relativistic chargedparticle beams with interstellar clouds, ap.j., 280, 550. [20] rose, w. k., beall, j. h., guillory, j., and kainer, s., 1987, radiation from relativistic beams interacting with interstellar gas, ap.j., 314, 95. doi:10.1086/165042 [21] saxton, c. j., bicknell, g. v., sutherland, r. s., and midgley s., 2005, interactions of jets with inhomogeneous cloudy media, mnras, 359, 781. doi:10.1111/j.1365-2966.2005.08962.x [22] scott, j. h., holman, g. d., ionson, j. a., and papadopoulos, k., 1980, the heating of gas in clusters of galaxies by relativistic electrons collective effects, ap.j., 239, 769. [23] su, m., and finkbeiner, d. p., 2012, evidence for gamma-ray jets in the milky way, ap.j., 753, 61. doi:10.1088/0004-637x/753/1/61 [24] zanni, c. et al., 2005, heating groups and clusters of galaxies: the role of agn jets, astron. astrophys., 429, 399. discussion manel perucho could you comment on the viscosity of the bow shock in your 3-d simulations? what is causing it? jim beall the viscosity is likely due to turbulence generated by the jet-ambient medium interaction and from internal shock structures within the jet and in the jet interface with the interstellar medium. gennady bisnovatyi-kogan was a magnetic field included in your numerical simulations? jim beall the vh-1 code we are using is a hydrodynamic code with no capability for the inclusion of magnetic fields. we can estimate the magnetic field strength by assuming equipartition with the gas pressure, but this is not entirely satisfactory. we want to explore the pluto code, which has an mhd capability, in the near future. 277 http://dx.doi.org/10.1086/155845 http://dx.doi.org/10.1007/978-94-009-0545-0_20 http://dx.doi.org/10.1016/0021-9991(84)90143-8 http://dx.doi.org/10.1086/322479 http://dx.doi.org/10.1016/s1387-6473(03)00096-4 http://dx.doi.org/10.1088/0004-6256/137/3/3718 http://dx.doi.org/10.1063/1.859385 http://dx.doi.org/10.1086/165042 http://dx.doi.org/10.1111/j.1365-2966.2005.08962.x http://dx.doi.org/10.1088/0004-637x/753/1/61 introduction scales of jet interactions with the ambient medium energy loss, energy deposition rate, and momentum transfer from plasma processes results of hydrodynamical calculations concluding remarks 240 acta polytechnica ctu proceedings 1(1): 240–245, 2014 240 doi: 10.14311/app.2014.01.0240 unveiling the nature of integral objects: a review pietro parisi1, on behalf of a large collaboration 1istituto di astrofisica e planetologia spaziali (inaf), via fosso del cavaliere 100, roma i-00133, italy corresponding author: p. parisi abstract since its launch in october 2002, the integral observatory has improved our knowledge of the hard x-ray sky above 20 kev, carrying out more than ten years of observations in the energy range from 5 kev to 8 mev. the most recently published integral/ibis surveys listed more than seven hundred sources in the 20-100 kev band. most of these objects are either active galaxies (agns) or x-ray binaries; a fraction of both classes is made of highly absorbed sources, often associated with dim optical counterparts. despite the big effort in the identification process, a large part of these ibis objects (∼25% of them) still remains unclassified. cross-correlation with archival catalogues and/or multiwaveband follow-up observations are of invaluable help to identify and properly classify this unknown objects, but only optical or ir spectroscopy with ground based telescopes in the northern and southern hemisphere can reveal the real nature of these objects. in this work we report on source types that we find among the unidentified objects in the most recent integral surveys. keywords: galaxies: seyfert x-ray binaries techniques: spectroscopic. 1 introduction one main objective of the integral mission is a regular survey of the entire sky in the hard x-ray band. this makes use of the unique imaging capability of the ibis instrument (ubertini et al. 2003) which allows the detection of sources at the mcrab level with a typical localization accuracy of a few arcmin. during the first 5 years of its life, the observatory concentrated mainly on a deep exposure of the galactic central radian, regular scans of the galactic plane, pointed observations of the vela region, and target of opportunity follow up observations. through the ibis imager, optimized for survey work with excellent imaging and spectroscopy capability, catalogs devoted either to the galactic plane scan (gps) or to the improvement of the extragalactic coverage were published. up to now, ibis detected more than 700 sources in the hard x-rays between 20 and 100 kev (bird et al. 2010, krivonos et al. 2010, 2012), and ∼ 30% had no obvious counterpart at other wavelengths and therefore cannot yet be associated with any known class of high-energy emitting sources. comparisons with catalogues or surveys at other frequencies (especially soft x-rays, optical, infrared and radio) are of invaluable help in reducing the localization uncertainty of these ibis sources from arcminutes down to a few (< 10) arcseconds, thus making the search for their optical or ir counterparts much easier. once this is done, then spectroscopic follow up observations provide a classification of the source, a confirmation of the proposed association and the study of that source individually or at population level. thanks to the above method, many new hard x-ray sources have been studied for the first time, including new classes of galactic objects, such as absorbed high mass x-ray binaries (hmxb see walter 2007), supergiant fast x-ray transients (sfxt, e.g., sguera et al. 2005, 2006; leyder et al. 2007), magnetic cataclysmic variables (cvs; barlow et al. 2006; bonnet-bidaud et al. 2007; landi et al. 2008) and symbiotic x-ray binaries (syxbs); in the extragalactic sky, a higher percentage of absorbed active galactic nuclei (agns) were reported compared to softer (2-10 kev) surveys, including a few new nearby compton thick objects (malizia et al. 2009), systematically pinpointing, for the first time, extragalactic sources in the so-called ‘zone of avoidance’, which hampers observations in soft x-rays and optical along the galactic plane due to the presence of gas and dust. all of these objects were classified by means of intensive optical and ir spectroscopic campaigns at various telescopes located worldwide. 2 surveys through hard x-ray surveys we can obtain all-sky maps of the celestial high-energy emission and study catalogues of sources unbiased in term of absorption and which are capable of producing non thermal emission 240 http://dx.doi.org/10.14311/app.2014.01.0240 unveiling the nature of integral objects: a review processes, or being the site of the most extreme astrophysical phenomena observed in the universe. the most recent integral/ibis surveys are that of bird et al. (2010) and krivonos et al. (2010, 2012): • the 4th ibis survey (bird et al. 2010) collected data from november 2002 to october 2008 for ∼80 ms of exposure in the 20-100 kev energy range and produced a catalogue of 723 sources (29% unidentified) ; • the ibis 7-years all-sky hard x-ray survey (krivonos et al. 2010) collected data from december 2002 to july 2009 for ∼80 ms of exposure in the 17-60 kev energy range and found 521 sources (12% unidentified); • the 9-year galactic hard x-ray survey (krivonos et al. 2012) collected data from december 2002 to january 2011 for 132 ms of exposure in the 17-60 kev energy range and reported 402 sources (9% unidentified). 3 identification method in the second semester of 2004 we started an optical spectroscopy campaign performed at ground based telescopes of the northern and southern hemisphere to identify unknown hard x-ray sources detected by integral (masetti et al. 2004, 2006a,b,c,d, 2007, 2008a,b, 2009, 2010, 2012, 2013), and we selected unidentified or unclassified hard x-ray sources that contain, within the ibis 90% confidence level error box, a single bright x-ray object detected either in the rosat all-sky surveys (voges et al. 1999, 2000), or in the slew survey (saxton et al. 2008) and/or in the serendipitous source catalog (watson et al. 2009) of xmm-netwon, or having pointed observations either from chandra, swift/xrt or xmm-netwon satellites. this approach was proven by stephen et al. (2006) to be very effective in associating, with a high degree of probability, ibis sources with a softer x-ray counterpart, in turn drastically reducing the positional error circles to better than a few arcsec in radius, making the search area smaller by a factor of 104. taking into account the hard x-ray characteristics and the global properties of the sources, such as the position in the sky, the 20-100 kev light curve and the broad band spectrum, we can have some clues about their nature. after this first selection, we chose among these objects those that had, within their refined 90% confidence level soft x-ray error boxes, a single possible optical counterpart with magnitude r < 20 in the dss-ii-red survey, so that optical or ir spectroscopy could be obtained with reasonable signal-tonoise ratio using medium-sized telescopes (i.e. with diameter up to 4 meters; masetti et al. 2004, 2006a,b,c,d, 2007, 2008a,b, 2009, 2010, 2012, 2013). 4 galactic sources many important results from the integral mission have been obtained in the observations of the galactic sources. specifically, they have shown the existence of a new class of heavily absorbed x-ray binaries and of the sfxts, doubling the number of known hmxbs and allowed the detection of a substantial number of new magnetic cvs. 4.1 cvs cvs are close binary systems consisting of a late-type (i.e. red dwarf) star trasferring material onto a white dwarf (wd) via roche lobe overflow. those which have a magnetic field are called magnetic cvs and fall into two categories: • polar cvs have a strong magnetic field (b > 107 g); their accretion does not occur via accretion disc and the accreting material is channeled by the magnetic field along its lines and falls on the magnetic poles of the wd; • intermediate polars (warner 1995), instead, have a weaker magnetic field (b ∼ 106 − 107 g) that truncates the accretion disc in the inner region close to the magnetosphere, resulting in an accretion curtain, where the accreting material follows the magnetic field lines down to the wd poles. if the wd is not magnetic the accreting material flows towards the wd through an accretion disc. up to now the cvs detected by integral are 35 (∼ 80% are magnetic, mostly intermediate polars), and 29 of these have been identified through optical or nir spectroscopy. this is an important result, if we consider that before integral very few cvs have been detected at high energies. fig. 1 shows the cvs distribution in the sky (white filled circles superimposed on an ibis 20-100 kev image). figure 1: distribution of cvs (white filled circles) in the 20-100 kev sky imaged by ibis. 241 pietro parisi 4.2 low mass x-ray binaries low mass x-ray binaries (lmxbs) are systems consisting of an accreting compact object (neutron star or black hole) and a low-mass (< 1 m�) main-sequence or slightly evolved late-type star. the low mass companion fills and overflows its roche lobe, therefore accretion of matter always occurs through the formation of an accretion disc. up to now ∼ 100 lmxbs have been detected by integral, and only 15 have been classified through optical or nir spectroscopy, this means that almost all the lmxbs have been detected and very few are found among unidentified integral sources. in fig. 2 we report the lmxbs distribution (white filled circles superimposed on an ibis 20-100 kev image): it is clear that they lie in the galaxy bulge or in globular clusters, where we the majority of old stellar populations are located. figure 2: distribution of lmxbs (white filled circles) in the 20-100 kev sky imaged by ibis. they are segregated in the galactic bulge and in globular clusters. 4.2.1 symbiotic x-ray binaries three of the lmxbs spectroscopically identified among unidentified objects are symbiotic x-ray stars (syxbs). they are part of a small subclass of lmxbs in which the compact object (generally a neutron star), receives matter from a red giant rather than from a late-type companion star on the main sequence. compared to lmxbs they show an optical continuum typical of a red giant of m spectral type, with balmer series generally in absorption (of the 7 syxbs known only one, gx 1+4, shows hα in emission, because the donor star is a red supergiant star that fills its roche lobe, accreting matter via accretion disc) and with the presence of absorption bands, and matter is accreted via stellar wind from the donor star. 4.3 hmxbs the galactic plane scans performed with integral revealed a wealth of new hmxbs. these binary systems are composed of a compact object (neutron star or black hole) orbiting and accreting matter from a luminous early spectral type ob high mass (> 10 m�) companion star. hmxbs have different ways to accrete matter: • via roche lobe overflow, but we know only very few cases of this type; • star with a circumstellar disc: the compact objects with a wide eccentric orbit crosses the decretion disc produced by a rapidly rotating be iii/iv/v star, producing accretion through stellar wind; • a massive star (supergiant i/ii star) ejects a fast and dense radially outflowing wind, and the compact object directly accretes from it. up to now ∼ 90 hmxbs have been detected by integral; of these, 45 were identified through optical or nir spectroscopy. this class of objects is distributed along the galactic plane (see fig.3). if we also consider the hmxbs distances, we note that they closely trace the underlying distribution of the massive starforming regions that are expected to produce the progenitor stars of hmxbs (bodaghee et al. 2012). figure 3: distribution of hmxbs (white filled circles) in the 20-100 kev sky imaged by ibis. they lie along the galactic plane. 4.3.1 sfxts this is a new class of transient hmxbs discovered by integral. the sfxts host a massive blue supergiant star (ob) and a compact object, mainly a neutron star. they have fast x-ray flares, from few hours to few days duration, with luminosity of ∼1036-1037 erg s−1; they also have short duty cycles (tinflare/ttotal=0.05% 3%) the accretion mechanism is not clear yet. there are different scenarios, such as the clumpy wind (negueruela et al. 2008, ducci et al. 2009) or the centrifugal/magnetic barrier (bozzo et al. 2008). the clumpy wind scenario has two possible configurations: a neutron star orbiting a supergiant star on a circular orbit, or on an eccentric orbit, accreting 242 unveiling the nature of integral objects: a review from the clumpy stellar wind of the supergiant. in the centrifugal/magnetic barrier scenario, a magnetic barrier or a centrifugal barrier can sets in, and according to the spin period and the strength of the magnetic field of the neutron star, it can cause or not the inhibition of accretion. up to now the sfxts optically or nir identified are ∼10 and other ∼10 are candidates. 5 extragalactic objects the integral satellite has been able to obtain also important results in the field of extragalactic objects. observations are giving fundamental insights into the study of agns located in the zone of avoidance along the galactic plane (see fig. 4 for the agns distribution in the sky) but also positioned across the whole sky. figure 4: distribution of agns (white filled circles) in the 20-100 kev sky imaged by ibis. this is an ibis 20-100 kev image. 5.1 seyfert galaxies the most common extragalactic objects detected by integral are seyfert galaxies; their luminosity ranges from 1042 to 1045 erg s−1, they are located in the nearby universe (z < 0.5) and have an optical spectrum characterized by emission lines. all the seyferts show narrow high ionization emission lines, such as [oiii] or [nii] forbidden emission lines. some show broad permitted lines in emission (generally balmer series), suggesting the presence of a dense and fast moving gas: we call these seyfert 1 galaxies. in these objects, according to the unified model, our line of sight intercepts both broad line regions and narrow line regions. those agns with only narrow emission lines are instead named seyfert 2 galaxies. according to the unified model, in these sources our line of sight intercepts only the narrow line region due to the presence of a torus which hides the broad line region. these gas and dusty regions are photoionized by the central engine; blrs are confined to sub-pc scales around an accretion disk, producing kinematically broadened emission lines with typical velocities of 103–104 km s−1, nlrs are characterized by narrow-lines with typical velocities of 102–103 km s−1 and can span over kpc scales, which are comparable to the size of the bulge or even the entire galaxy. up to now 148 agns detected by integral have been identified through optical and nir spectroscopy. in particular, 68 agns are seyfert 1, while 55 are seyfert 2 galaxies. 5.1.1 narrow-line seyfert 1 these galaxies are peculiar seyfert 1 agns (osterbrock & pogge 1985) with a full width at half maximum (fwhm) of the hβ emission line smaller than 2000 km s−1, with permitted lines which are only slightly broader than the forbidden ones, with a [oiii]5007/hβ ratio < 3, and finally with evident feii and other highionization emission-line complexes (e.g. see fig. 5). a few nls1 galaxies have been discovered so far by integral and recognized as such by optical spectroscopy. figure 5: optical spectrum (not corrected for the intervening galactic absorption) of a typical narrow-line seyfert 1. the main spectral features are labeled. 5.2 low ionization nuclear emission-line regions low ionization nuclear emission-line regions (liners; heckman 1980) are peculiar agns with a level of activity much smaller than that in classical agns and in which some low-ionization lines ([oii]3723, [oi]6300, and [nii]6584) are stronger than in typical seyfert 2 galaxies; the permitted emission-line luminosities are weak; and the emission-line widths are comparable with those of type 2 agns (see fig. 6). according to heckman (1980), liners have [oii]3723 > [oiii]5007 and [oi]6300 > 1/3 [oiii]5007, and often [nii]6584/hα > 0.6. 243 pietro parisi figure 6: optical spectrum (not corrected for the intervening galactic absorption) of a typical liner. the main spectral features are labeled. 5.3 x-ray bright optically normal galaxies some integral objects show a continuum typical of a normal galaxy, dominated by absorption lines due to star forming regions, they are x-ray bright, optically normal galaxies (xbongs; comastri et al. 2002), that is, x-ray bright galactic nuclei with no emission lines in their optical spectra (e.g. fig. 7). figure 7: optical spectrum (not corrected for the intervening galactic absorption) of a typical xbong. the main spectral features are labeled. 5.4 blazars using medium-sized telescopes (i.e. tng, eso), we are also able to identify and classify sources at high redshifts (> 0.6) detected by integral. they are blazars, distant and powerful agns which are oriented in such a way that a jet expelled from the central black hole is directed at small angles with respect to the observers line of sight. an example is igr j12319-0749 a powerful blazar at z = 3.12 (masetti et al. 2012), the farthest opticallyidentified object of any integral survey and the second furthest of all objects detected by integral. 6 summary and conclusions up to now 273 integral objects have been identified and classified through optical and nir spectroscopy using ground-based telescopes of the northern and southern hemisphere. of these objects, 62% are agns, 36% are x-ray binaries and the remaining 2% are chromospherically active stars. going into details, of the 62% agns, 27.8% are seyfert 1 galaxies, 22% are seyfert 2 agns, 8.5% qsos, 2.9% xbongs and 1.1% are other sources. among the galactic sources (36%), we found that 18% are hmxbs, 12.5% are cvs and 5.5% are lmxbs. the ibis surveys secured the detection of extragalactic sources in the so-called zone of avoidance, which hampers observations in soft x-rays along the galactic plane due to the presence of gas and dust. moreover, these surveys are expanding our knowledge about galactic x-ray binaries, by showing the existence of a new class of heavily absorbed supergiant massive xray binaries (first suggested by revnivtsev et al. 2003), by allowing the discovery and the study of supergiant fast x-ray transients (e.g., sguera et al. 2005, 2006; leyder et al. 2007), by doubling the number of known hmxbs (see walter 2007), and by detecting a substantial number of new magnetic cvs. indeed, the new ibis catalogue (in progress) will offer new spectral and timing information on newly detected sources and an insight on peculiar ones, giving us the unique opportunity to discover new hmxbs and understand the differences among them. references [1] barlow, e. j., knigge, c., bird, a. j., et al. 2006, mnras, 372, 224 doi:10.1111/j.1365-2966.2006.10836.x [2] bird, a. j., bazzano, a., bassani, l., et al. 2010, apjs, 186, 1 doi:10.1088/0067-0049/186/1/1 [3] bodaghee, a., tomsick, j.a., rodriguez, j., et al. 2012, apj, 744, 108 doi:10.1088/0004-637x/744/2/108 [4] bonnet-bidaud, j. m., de martino, d., falanga, m., mouchet, m., & masetti, n., 2007, a&a, 473, 185 244 http://dx.doi.org/10.1111/j.1365-2966.2006.10836.x http://dx.doi.org/10.1088/0067-0049/186/1/1 http://dx.doi.org/10.1088/0004-637x/744/2/108 unveiling the nature of integral objects: a review [5] bozzo, e., falanga, m. & stella, l., 2008, apj, 683, 1031 doi:10.1086/589990 [6] comastri, a., mignoli, m., ciliegi, p., et al. 2002, apj, 571, 771 doi:10.1086/340016 [7] ducci, l., sidoli, l., mereghetti, s., et al. 2009, mnras, 398, 2152 doi:10.1111/j.1365-2966.2009.15265.x [8] heckman, t. m. 1980, a&a, 87, 152 [9] krivonos, r., tsygankov, s., revnivtsev, m., et al. 2010, a&a, 523, a61 [10] krivonos, r., tsygankov, s., lutovinov, a., et al. 2012, a&a, 545, [11] landi, r., bassani, l., dean, a. j., et al. 2008, mnras, 392, 630 doi:10.1111/j.1365-2966.2008.14086.x [12] malizia, a., stephen, j.b., bassani, l., et al. 2009, mnras, 399, 944 doi:10.1111/j.1365-2966.2009.15330.x [13] masetti, n., palazzi, e., bassani, et al. 2004, a&a, 426, l41 [14] masetti, n., mason, e., bassani, l., et al. 2006a, a&a, 448, 547 [15] masetti, n., pretorius, m.l., palazzi, e., et al. 2006b, a&a, 449, 1139 [16] masetti, n., bassani, l., bazzano, a., et al. 2006c, a&a, 455, 11 [17] masetti, n., morelli, l., palazzi, e., et al. 2006d, a&a, 459, 21 [18] masetti, n., landi, r., pretorius, m.l., et al. 2007, a&a, 470, 331 [19] masetti, n., mason, e., morelli, l., et al. 2008a, a&a, 482, 113 [20] masetti, n., mason, e., landi, r., et al. 2008b, a&a, 480, 715 [21] masetti, n., parisi, p., palazzi, e., et al. 2009, a&a, 495, 121 [22] masetti, n., parisi, p., palazzi, e., et al. 2010, a&a, 519, a96 [23] masetti, n., parisi, p., jiménez-bailón, e., et al. 2012, a&a, 538, a123 [24] masetti, n., parisi, p., palazzi, e., et al. 2013, a&a, in press. [25] negueruela, i., torrejon, j. m., reig, p., ribo, m., & smith, d. m. 2008, in a population explosion: the nature & evolution of x-ray binaries in diverse environments. aip conference proceedings, volume 1010, 252 doi:10.1063/1.2945052 [26] osterbrock, d.e., & pogge, r.w. 1985, apj, 297, 166 doi:10.1086/163513 [27] revnivtsev, m. g., sazonov, s. y., gilfanov, m. r., & sunyaev, r. a. 2003, astron. lett., 29, 587 doi:10.1134/1.1607496 [28] saxton r.d., read, a.m., esquej, p., et al. 2008, a&a, 480, 611 [29] sguera, v., barlow, e. j., bird, a.j., et al. 2005, a&a, 444, 221 [30] sguera, v., bazzano, a., bird, a.j., et al. 2006, apj, 646, 452 doi:10.1086/504827 [31] stephen, j.b., bassani, l., malizia, a., et al. 2006, a&a, 445, 869 [32] ubertini, p., lebrun, f., di cocco, g., et al. 2003, a&a, 411, l131 [33] voges, w., aschenbach, b., boller, t., et al. 1999, a&a, 349, 389 [34] voges, w., aschenbach, b., boller, t., et al. 2000, iau circ. 7432 [35] walter, r. 2007, ap. space sci., 309, 5 doi:10.1007/s10509-007-9477-9 [36] warner, b. 1995, cataclysmic variable stars (cambridge: cambridge university press) doi:10.1017/cbo9780511586491 [37] watson, m.g., schroder, a.c., fyfe, d., et al. 2009, a&a, 493, 339 discussion pieter meintjes: the novalike cataclysmic variable ae aquarii has been detected by suzaku above 10 kev, showing evidence of non-thermal emission. has it been detected by integral as well? pietro parisi: the cataclysmic variable ae aquarii has not been observed by integral. 245 http://dx.doi.org/10.1086/589990 http://dx.doi.org/10.1086/340016 http://dx.doi.org/10.1111/j.1365-2966.2009.15265.x http://dx.doi.org/10.1111/j.1365-2966.2008.14086.x http://dx.doi.org/10.1111/j.1365-2966.2009.15330.x http://dx.doi.org/10.1063/1.2945052 http://dx.doi.org/10.1086/163513 http://dx.doi.org/10.1134/1.1607496 http://dx.doi.org/10.1086/504827 http://dx.doi.org/10.1007/s10509-007-9477-9 http://dx.doi.org/10.1017/cbo9780511586491 introduction surveys identification method galactic sources cvs low mass x-ray binaries symbiotic x-ray binaries hmxbs sfxts extragalactic objects seyfert galaxies narrow-line seyfert 1 low ionization nuclear emission-line regions x-ray bright optically normal galaxies blazars summary and conclusions 192 acta polytechnica ctu proceedings 2(1): 192–196, 2015 192 doi: 10.14311/app.2015.02.0192 the role of magnetic field for quiescence-outburst models in cvs s. de bianchi1,2, v. f. braga3, s. gaudenzi1 1dept. of physics, università degli studi di roma ”la sapienza”, piazzale a. moro 5, 00185 rome, italy 2école normale supérieure, cnrs-umr 8547, rue d’ulm 45, 75005 paris, france 3dept. of physics, università tor vergata, via della ricerca scientifica 1, 00133 rome, italy corresponding author: silvia.debianchi@uniroma1.it abstract in this paper we present the elementary assumptions of our research on the role of the magnetic field in modelling the quiescence-outbursts cycle in cataclysmic variables (cvs). the behaviour of the magnetic field is crucial not only to integrate the disk instability model (osaki 1974), but also to determine the cause and effect nexus among parameters affecting the behavior of complex systems. on the ground of our interpretation of the results emerging from the literature, we suggest that in models describing dne outbursts, such as the disk instability model, the secondary instability model (bath 1973) and the thermonuclear runaway model (mitrofanov 1978), the role of the magnetic field is at least twofold. on the one hand, it activates a specific dynamic pathway for the accreting matter by channelling it. on the other hand, it could be indirectly responsible for switching a particular outburst modality. in order to represent these two roles of the magnetic field, we need to integrate the disk instability model by looking at the global behaviour of the system under analysis. stochastic resonance in dynamo models, we believe, is a suitable candidate for accomplishing this task. we shall present the mhd model including this mechanism elsewhere. keywords: cataclysmic variables dwarf novae intermediate polars magnetic field disk instability model secondary instability model thermonuclear runaway model. 1 introduction in the last decades, methods and models used to describe dne outburst cycles achieved several results (bath & van paradijs 1983, cannizzo & mattei 1998, cannizzo 2012). however, there is no unique model that predicts these cycles (smith 2007), as well as their underlying mechanism. the limit-cycle model for dwarf nova outbursts is generally accepted, but there are still uncertainties about how to explain all the details, especially for the su uma stars (smith 2007). even if we know that the secondary stars in dne are magnetically active, yet it is unknown how they maintain a dynamo in the presence of tidal forces or whether there is differential rotation (smith 2007). crawford et al. (2008), for instance, when presenting their results of the detection of the first observed outburst of dw cnc, asked the question of what caused the outburst observed. dw cnc experienced a magnitude change of ∼4 mag showing a behavior similar to those systems whose outbursts are due to disk instability. however, with respect to pspin and porb, dw cnc shares characteristics of those systems whose outbursts are due to a mass transfer event. therefore, a decision for which mechanism was responsible for the first detected outburst could not be made. precisely in order to overcome this kind of problems of undecidability a deeper analysis of our models’ ability to explain dne outbursts is needed. 2 disk instability model and secondary instability model: some open questions each of our best available models, such as the disk instability model (dim), the secondary instability model (sim) and thermonuclear runaway (tnr) model, taken separately, shows weaknesses in predicting the outburst-quiescence cycles of many systems, such as v513 cas for outbursts during standstills (hameury & lasota 2014). dim is generally accepted as reproducing an explanation of dne outbursts and invokes an intrinsic modulation of the accretion rate in the disk. however, as seen in the case of dw cnc, the model does not satisfactorily account for all systems and suffers of some structural problems. the first one has been inherited by the disk model of shakura & sunyaev (1973) which was based on a constant value of α, a parameter that stores all unknown information of the complex friction processes. a constant α does not account for a reliable outburst amplitude (smak, 1984) and the pa192 http://dx.doi.org/10.14311/app.2015.02.0192 the role of magnetic field for quiescence-outburst models in cvs rameter appeared to be an ad hoc solution for strictly regular outbursts (bath 2004). studies concerning its variation or a turbulent α parameter aimed to the appropriate simulation of the disk instability by applying suitable corrections (latter & papaloizou 2012; penna et al. 2013; potter & balbus 2014). a second problem consists in the fact that dim does not appropriately model the observed fluctuations, due to the assumption of a constant value for mass transfer rate ṁ. third, dim shows weaknesses in predicting global changes in the disk structure, such as whether thin disk accretion can make a transition to advection-dominated flow (narayan & yi 1994, 1995a, 1995b) or how an accretion disk creates and powers jets at its center (king 2012). as an alternative to dim, sim assumes a mechanism triggering the outbursts based on mass transfer modulation from the secondary star (bath et al. 1974). by combining dynamical instabilities of the secondary, time evolution of the accretion disk together with thermonuclear burning due to accreted material, sim produces reliable results: outbursts during standstills in z cam systems can be explained only by appealing to instabilities in the flow from the secondary (hameury & lasota 2014). it seems then that the sim is decisive in predicting cvs like z cam variables and, possibly, su uma variables. nevertheless, its application has not been always straightforward. osaki (1985), for instance, explained the superoutbursts of su uma variables by using sim, but later denied this possibility in favor of a thermal-tidal-instability (tti) model (osaki 1989). finally, smak (1996) suggested a hybrid ttisim model, which is still debated today (osaki & kato 2013). 3 magnetic field and outburst modality in order to enrich the theoretical scenarios in dealing with dne outburst-quiescence cycle, the (tnr) model shows interesting implications, even if the improvement of the tnr model launched by shara (1982) never brought to a decisive point. tnr introduces a time parameter describing the recurrence of outbursts accounting for both classical novae (cne) and dne outbursts and their differences (shara 1982) and also accounts for dnos and x-ray flux variations concerning some representative systems like u gem, ss cyg, ex hya, z cam, cy cnc, ah her, cn ori, kt per. more importantly, tnr can be associated to one of the processes that causes nitrogen to carbon enhancement observed on the surface of the wd in vw hyi and u gem (sion 2014). in tnr models, outbursts are the product of thermonuclear burning onto the wd surface (mitrofanov 1978; 1980). in the presence of a magnetic field of the order of 3 · 106 − 3 · 107 g on the surface of the degenerate component, a local accumulation of hydrogen could generate, through thermal instability, a thermonuclear burning responsible of the outburst. for this process to occur, an intense magnetic field 106−108 g is required; primaries with such fields do exist in intermediate polars (ips) and polars that owe their properties to the very primary magnetic field (patterson 1994). therefore, the magnetic field strength on the wd surface allows us to distinguish novae, dwarf novae and novalike stars. in exploring models that go beyond the use of tnr alone, livio (1983) emphasized the importance of the magnetic fields in cvs (livio & verbunt 1988; meyer-hofmeister et al. 1996). as livio (1983) shows, nova explosions can be inhibited if the magnetic field strength is over a certain limit, explaining in this way why active novae are absent among polars while novae can be found among the ips. figure 1: upper limits on the magnetic field strength for which nova outbursts can occur (provided that matter is confined to polar caps) as a function of the accretion rate, for various white dwarf masses. reproduced from livio, m., 1983, astronomy and astrophysics, 121, l7, with permission from astronomy and astrophysics, c©eso. livio’s approach allows one to determine upper limits on the magnetic field strength for which nova outbursts can occur when polar caps are present. these upper limits can be represented as a function of the accretion rate and for different wd masses (see figure 1). even if a magnetic field more intense than a certain critical value could inhibit nova explosions (livio 1983), a magnetic field is strictly correlated, on an intensity-dependent scale, to different modes of accretion and thermonuclear runaways. these modes in193 s. de bianchi, v. f. braga, s. gaudenzi clude the channeling of the accreted matter into magnetic polar caps, the alteration of radiative and conductive opacities, interference with the development of convection that likely is also present in non magnetic cvs. furthermore, the magnetic field of the wd itself may affect the nova outbursts, e.g. by enhancing mass loss in the equatorial plane (livio et al. 1988; prialnik & livio 1995). this reveals how the magnetic field activates different and specific dynamic pathways. in other words, to focus on the role of a variable magnetic field and its effects can be helpful for identifying common features of both magnetic and non-magnetic cvs, as well as for obtaining crucial information on the magnetic field responsible for switching a particular outburst modality. indeed, according to livio (1983), the pressure at the base of the accreted matter is the physical parameter that determines the outcomes of a tnr: pb = gmw d r2w d ∆macc acap (1) when pb exceeds a critical value of some 1019dyne/cm2, ignition occurs. it should be noted that, for higher cno abundances, the critical pressure for the occurrence of an outburst is lower. this fact leads us to the investigation of cno abundances and of the presence of material accumulating during quiescence, in order to speculate on the triggering of an outburst. it means that a function of “prediction” and not of mere “description” can be added to the model. figure 2: e(b-v) varies with both the orbital phase and the percentile quiescence time (t%). 2175 å absorption bump is believed to be caused by c60. e(b-v) takes the highest value at the “heart” of the quiescence. c60 molecules could be reasonably supposed to be affecting the modulation of time elapsing between the outbursts. reproduced from gaudenzi et al. (2011), astronomy and astrophysics 525, a147 with permission from astronomy and astrophysics, c©eso. more importantly, the analysis of uv spectra revealed the presence of fullerenes as an intrinsic source of reddening in ss cyg (gaudenzi et al. 2011, see figure 2). density gradients in the disk may influence the accumulation of molecules in specific sites of the disk itself where they are accreted and/or ejected during the quiescence-outburst cycle. along with accretion/ejection mechanisms, a major role in determining the structure and the stability of the disk might be played by the solid-gaseous phase transitions of fullerenes (gaudenzi et al. 2012) at 5855 k (hussien et al. 2008). on the ground of the hypothesis that the quiescence-outburst phase transition is the effect of a stochastic resonance mechanism, the presence of material accumulated in the disk could trigger wave amplification within the disk even during quiescence. this would introduce in our modelling of the disk instability an amplification of stochastic resonance induced by turbulent fluctuations (benzi & pinton 2011). 4 magnetic field and accretion behaviour seminal numerical studies of the influences of a magnetic field on the flow structure were performed in the early 1990s either in the frame of simplified models (king 1993; wynn & king 1995; wynn et al. 1997; king & wynn 1999; norton et al. 2004; ikhsanov et al. 2004; norton et al. 2008) or in a limited region of the stellar magnetosphere (koldoba et al. 2002; romanova et al. 2003, 2004). in the last few years, there have been attempts to develop a comprehensive 3d numerical model to calculate the flow structure in close binaries (zhilkin & bisikalo 2009, 2010; bisikalo & zhilkin 2012). according to our interpretation of the results of 3d mhd simulation of bisikalo & zhilkin (2012), it emerges that the value of the magnetic induction on the surface of the accreting star activates the dynamic pathway and, in doing so, distinguishes two different modalities of accretion. in the numerical model, bisikalo & zhilkin (2012) take into account radiative heating and cooling,as wells as diffusion of the magnetic field due to dissipation of currents in turbulent vortexes, magnetic buoyancy, and wave mhd turbulence. the interesting result consists in that if the magnetic field induction grows, the cross-section of the stream decreases and the accretion rate decreases as well, thereby influencing density and pressure, and thus affecting the possibility of triggering an outburst. we interpret the nonmonotonic variation of the magnetic field behaviour and its consequences reported in (bisikalo & zhilkin 2012) as stochastic oscillations of b. in particular, magnetic field values around 106g could originate phase transitions between polar and ip. based on our previous work (gaudenzi et al. 2012), we suggest that the dynamic pathway of accretion can account for instabilities of many cvs and discloses the possibility of acquiring an ability to predict their behaviour as transition ob194 the role of magnetic field for quiescence-outburst models in cvs jects. the physics behind this process concerns the amplification of stochastic resonance induced by turbulent fluctuations, i.e. the amplitude of the external periodic perturbation needed for stochastic resonance to occur is much smaller than the one estimated by the equilibrium probability distribution of the unperturbed system (benzi & pinton 2011). 5 discussion and conclusion in order to enrich the theoretical scenario we stressed the relevance of the accretion process that may change the surface magnetic field of an accreting wd significantly (cumming 2002), and, in agreement with (livio 1983), that the magnetic field of the wd itself may affect the nova outburst. we also remarked that the accumulation of some type of molecules, such as c60 molecules within the disk is important to understand the behavior of dne in quiescence and that the cno abundances also depend on the presence of a magnetic field (even of a weak one). if we want to predict the behavior of a complex system, such as dne outbursts cycles, neither dim nor any other model taken alone is sufficient. in fact, dim explains why thermal instability leads to outbursts, but it does not provide exact information at the level of the global system in the quiescent phase, e.g. whether there are transition from a state to another, and how to determine time-dependent α variations. in order to obtain more information about the dne outburst-quiescence cycle, we shall work on a model that is able to predict future states of these systems, as well as to describe the present state, without appealing to ad hoc assumptions. as previously stated, local, small-scale turbulences, may likely represent the physical outcome of stochastic oscillations of the magnetic field. among the other mechanisms, such as a strong differential rotation and vertical density gradients (pudritz 1981b), turbulences can be responsible for the generation of a large-scale magnetic field (pudritz 1981a). therefore, the development of mean field dynamo theory is crucial in order to overcome some difficulties related to current models in accounting for the role of the magnetic field generated in a turbulent medium. this theory, if appropriately modified, could allow us to calculate the overall structure of the global field without any detailed knowledge of the small-scale turbulence. even if this echoes earlier works by shakura & sunyaev (1973) there is a characteristic of the mean field dynamo theory that makes it different from earlier theories. it builds up a large-scale magnetic field trying to achieve a global redistribution of the angular momentum, whereas other models assume a smallscale field and a local viscosity. therefore, in order to gain knowledge on the underlying mechanism producing outbursts, such as the one presented in dw cnc, we will 1) simulate the interaction among thermonuclear runaway and magnetic field leading to a specific modification of the α viscosity parameter to be integrated in the model (switching the outburst modality); 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[54] zhilkin, a.g., bisikalo, d.v.: 2010, adspr, 45, 2420. 196 http://dx.doi.org/10.1093/mnras/sts185 http://dx.doi.org/10.1093/mnras/stu519 http://dx.doi.org/10.1086/377514 http://dx.doi.org/10.1086/421867 http://dx.doi.org/10.1086/160376 http://dx.doi.org/10.1086/131295 http://dx.doi.org/10.1080/00107510601181175 http://dx.doi.org/10.1093/mnras/286.2.436 introduction disk instability model and secondary instability model: some open questions magnetic field and outburst modality magnetic field and accretion behaviour discussion and conclusion acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0012 acta polytechnica ctu proceedings 5:12–16, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app on operation of 740 metre long freight trains on czech ten-t railway network michal drábek∗, vít janoš, zdeněk michl department of logistics and management of transport, faculty of transportation sciences, ctu in prague, czech republic ∗corresponding author: xdrabek@fd.cvut.cz abstract. regulation (eu) no 1315/2013 defines actual scope of core and comprehensive ten-t network, including both networks for railway freight transport. for the core network, possibility to operate 740 m long freight trains is required. the aim of this paper is to analyse availability of appropriate overtaking tracks for 740 m long freight trains. due to etcs braking curves and odometry, such trains, after etcs implementation, will require 780-800 m long overtaking tracks. for practical reasons (e.g. bypass lines), whole czech railway ten-t network is analysed. the overtaking track, whose occupation means influence on scheduled traffic or threat to boarding passengers, are excluded. the data was collected from station schemes from collection of official requisites for 2015/16 timetable, issued by sždc, czech state infrastructure manager. most of appropriate tracks are over 800 m long, but their density in the network and in particular directions varies considerably. for freight traffic, gradient of the line is important, so in the resulting figure, there are marked significant peaks for particular lines as well. czech ten-t lines are further segmented on the basis of number of tracks and their traffic character. then, specific issues on overtaking or crossing of 740 m long freight trains are discussed. as a conclusion, for long-term development of czech ten-t lines, targeted investment is recommended not only for passenger railway, but also for freight railway. an attractive capacity offer for railway undertakings, which can stimulate freight traffic on european rail freight corridors, can be represented by network-bound periodic freight train paths with suitable long overtaking tracks outside bottlenecks. after the overtaking by passenger trains, a freight train should run without stop through large node station or a bottleneck area. before the sections with high gradients, coupling of additional locomotives should be connected with the overtaking process. next suitable overtaking tracks should be available behind every significant peak of the line. keywords: european freight corridor, long freight train, overtaking track, freight train path, peak . 1. introduction directive 2008/57/ec of the european parliament and of the council of 17 june 2008 on the interoperability of the rail system within the community [1] defines in annex ii infrastructure as a structural subsystem of the rail system, and traffic operation and management as a functional subsystem of the rail system. in the traffic operation and management subsystem, traffic planning is also included. regulation (eu) no 1315/2013 of the european parliament and of the council [2] defines actual scope of core and comprehensive ten-t network, including both networks for railway freight transport. for the core freight railway network, infrastructure requirements are defined. one of them is possibility to operate 740 m long freight trains. regulation (eu) no 913/2010 of the european parliament and of the council of 22 september 2010 concerning a european rail network for competitive freight [3] defined 9 initial rail freight corridors (rfcs) across the eu. according to the regulation, the corridors should provide guaranteed capacity for freight trains – "under good conditions in terms of commercial speed and journey times". the exact scope of each rfc are subject to negotiation by particular member states. three rfcs run directly through the czech republic: rfc 5 baltic-adriatic, rfc 7 orient and rfc 9 czech-slovak. moreover, negotiations about a branch of rfc 8 north seabaltic from germany to ústí nad labem and praha are proceeding. present, as well as planned or negotiated rfcs are displayed in the official presentation of czech ministry of transport by kušnír and ilík [4]. as a result, only few czech ten-t lines of comprehensive network are not included in any present rfc (or negotiated about being included in it). because of possible future changes of rfcs layout, and because of practical need for bypass lines, the whole czech railway ten-t network was included in the analysis. under the term "overtaking track" it is further understood any station track, which does not proceed any railway line in direct direction, i.e., it is not a main station track. for the sake of simplicity, the "not main" station tracks also in single-track lines, which serve mainly for crossing of trains, are referred to as "overtaking tracks". usable length of an overtaking track (further referred to only as "track length") is defined as a distance between last departure main signals, each one for the opposite direction. in the 12 http://dx.doi.org/10.14311/app.2016.5.0012 http://ojs.cvut.cz/ojs/index.php/app vol. 5/2016 on operation of 740 metre long freight trains case of group departure main signals (common for more parallel tracks), or absence of them, distance between last spots is considered, where rolling stock can be located, so that using of neighbouring track by another train is possible. the term "long trains" will be used for freight trains up to 740 m long. the term "long overtaking tracks" will be used for overtaking tracks at least 780 m long. for collection of data on length of overtaking tracks, railway station schemes from sždc (czech state infrastructure manager) collection of official requisites for 2015/16 timetable [5] were used. 2. methods 2.1. rules for determination of available overtaking tracks for 740 m long trains the basic question is, how long overtaking track is sufficient for a 740 m long train. for present czech signalling system and actual sždc regulations on traffic management, a track with additional 12 m length is sufficient. a 10 m distance is required between the locomotive and the departure main signal (in the train direction). the 2 m distance between an insulated gap and the departure main signal for the opposite direction is calculated. brejcha and čech [6] and binko [7] [8], however, state that for the future etcs implementation, a supplementary length of 40 to 60 m to the train length must be added, due to braking curve and inaccuracy of the etcs odometry. on the basis of the facts stated above, the authors have decided to analyse two categories of available long overtaking tracks: • 780 to 799 m long (in figure 1, separate presence of such track or tracks in particular direction is marked in yellow colour) • 800 m long and longer (in figure 1, presence of at least one such track in particular direction is marked in orange colour) presence of the tracks of both categories is marked in orange colour. presence of at least one long overtaking track in particular direction (on the train’s right, which is valid also for single-track lines or sections) is marked by obtuse-angled triangle of particular colour. its obtuse angle lays in direction, which is this track preferably used for. on czech railway network, there is only right-hand operation nowadays. in some recently modernized stations, an amended czech technical standard tnž 34 2620 railway station and line signalling and interlocking systems [9] was implemented. for the through run routes with allowed speed over 120 km/h, if there is no flank protection from the overtaking track and this track is occupied, only "restricted run route" (vco) with maximum speed 100 km/h is allowed. this mandatory function of interlocking systems obviously leads to longer runtime for the fast train than scheduled. so, overtaking of freight trains on vco-affected tracks is not desirable, but not critical. vco-affected long overtaking tracks are marked by pink edge in figure 1. for possible overtaking (or eventual crossing on single-track lines), only such overtaking tracks were considered, that did not lay between any main station track and waiting room (or entrance/exit of the station) in the case of at-grade access to platform (neither footbridge nor underpass). also such overtaking tracks were not considered that were (or were supposed to be, based on actual timetable [5]) occupied by passenger trains to be overtaken, or turned around. in stations with sidings, an expert estimation of need of station tracks for the freight trains that start or terminate here, was carried out. however, even such "occupied" or "reserved" long overtaking tracks were recorded (separately), if they were long enough. but they are neither displayed if figure 1 nor further discussed in this paper. in the case of a station with more yards (even a passenger and a freight yard), or with station tracks divided in the middle by switch region, an individual approach was chosen. if two subsequent tracks can be used by a long freight train without any impact on regular traffic, overall usable length of the resulting track was recorded. the czech ten-t lines will be further segmented into double-track lines, partially double-track lines and single-track lines. a triple-track line poříčany – praha-libeň and rather short triple-track sections bílina – odb české zlatníky and hranice na moravě – drahotuše will be added to double-track lines. the double-track group comprises vast majority of czech ten-t railway network. these groups can be further divided into subgroups on the basis of their traffic character. the presence of absence of appropriate overtaking tracks will be discussed by particular segments of the lines. 2.2. where to stop for overtaking – a soft decision process it is evident that any stop of a freight train in a bottleneck area leads to lower capacity utilization. so, overtakings should preferably be planned outside the bottlenecks. the following attributes of overtaking stations influence the quality of overtaking: • gradient on following line section (acceleration up the hill lengthens runtime of a freight train significantly and increases consumption of traction energy) • maximum allowed arrival speed on overtaking track (from the infrastructure viewpoint – it usually depends on arrival turnouts) • gradient on previous line section (braking downhill increases wear of brakes) • maximum allowable departure speed from overtaking track (usually depends on departure turnouts) • usable length of the longest overtaking track in particular direction 13 m. drabek, v. janos, z. michl acta polytechnica ctu proceedings in the authors’ opinion, it is very difficult (if not impossible) to set an exact order of priorities of the attributes listed above. each line, each passenger timetable and each set of freight trains together make a unique combination of specific conditions. so, a choice of the most suitable station for overtaking is always a soft decision process [10]. because of importance of gradients for freight railway, significant line peaks are also considered and marked in figure 1. 3. results and discussion czech ten-t lines are further segmented on the basis of number of tracks and their traffic character. presence or absence of long overtaking tracks is commented for each line group. for přerov and bohumín node stations, data on track length and on position of main departure signals were not available. some interesting findings occurred. firstly, during the analysis there emerged few theoretically suitable overtaking tracks which were approximately 770 m long. more precise calculations can determine the exact minimum practically usable track length for the long trains after implementation of etcs. secondly, most of the suitable overtaking tracks are over 800 m long. unfortunately, their density in the network varies considerably. 3.1. double-track (or triple-track) lines busy national mainlines with mixed traffic: • praha – česká třebová – olomouc – přerov – ostrava – bohumín (– chalupki) – dětmarovice/petrovice u karviné (– katowice) – třinec – bystřice busy mainlines with significant (or prevailing) freight traffic: • praha – děčín (– dresden), • lysá nad labem – děčín prostřední žleb, • ústí nad labem – úpořiny – bílina, • přerov – břeclav (– vienna/bratislava), • bystřice – mosty u jablunkova (– žilina). busy suburban sections with mixed traffic: • praha – kralupy nad vltavou, • praha – lysá nad labem – nymburk – kolín, • praha – benešov u prahy, • praha – beroun, • brno – tišnov, • brno – skalice nad svitavou, • brno – břeclav. middle-busy mainlines: • ústí nad labem – teplice – bílina – most – chomutov – karlovy vary – cheb (– marktredwitz), • plzeň – beroun, • kolín – kutná hora (station with both ac and dc traction power supply systems) – havlíčkův brod – tišnov, • česká třebová – skalice nad svitavou, • brno – blažovice, • hranice na moravě – horní lideč (– žilina). due to practically exhausted capacity of the line praha – olomouc – ostrava – bystřice, freight trains should be able to run homogeneous train paths with a long-distance passenger train, so their maximum speed should be 120 km/h, in the worst case 100 km/h. the locomotives should be enough powered to accelerate the train to such speed before it has to be overtaken by next fast train. even if such conditions are fulfilled, there are hardly any available train paths in morning and afternoon peaks between praha and česká třebová. because of busy traffic, the actual number of long overtaking tracks is not sufficient at all. moreover, some ones are located in kolín and pardubice, the busiest nodes within the section mentioned above. for a freight train it is extremely complicate to find a train path that enables conflict-free departure from such busy nodes with level crossing of trains in various directions (e.g. fast train praha – havlíčkův brod – brno with non-stop express train ostrava – pardubice – praha). the section kolín – lysá nad labem is busy by suburban traffic, but enables homogeneous train paths even to slower freight trains (e.g. 90 km/h). however, any stop of a freight train in this section is highly undesirable, because the trains run in 5-8 min wide time slots after each other. behind lysá nad labem, passenger traffic is considerably lower, but freight trains mostly do not need to stop because of train path homogeneity. in any case, the long trains should be scheduled early enough before a fast train, so that they could not delay them. most freight trains need rear-end or head-end assistance in line sections with high gradients, e.g. praha – strančice, český těšín – čadca, kutná hora – žďár nad sázavou, brno – žďár nad sázavou, hranice na moravě – horní lideč etc. for the long trains, 800 m long or longer overtaking tracks are desirable, because of length of the additional locomotive. it is appropriate to combine coupling and uncoupling of this locomotive with the overtaking process. due to speed profile of the line kralupy – děčín – dresden, even 100 km/h fast freight trains can run in homogeneous train paths with non-stop eurocity trains. there is lack of long overtaking tracks in the direction from praha. practically, a non-stop train path from praha-holešovice to děčín is required. on another lines of this sub-category, there is mostly almost no reason for the overtaking of freight trains, if the gradients are low. in busy suburban sections with mixed traffic, best capacity utilization means no stop of a freight train, which requires speed bundling (homogeneous train paths) either with suburban or with fast trains. 14 vol. 5/2016 on operation of 740 metre long freight trains figure 1. presence of available long overtaking tracks in czech ten-t railway network. map source: [11] 3.2. partially double-track lines (middle-busy mainlines) • plzeň – české budějovice (with double-track sections nepomuk – horažďovice předměstí and číčenice – zliv), • benešov u prahy – tábor – české budějovice (mostly double-track line, with remaining several single-track sections). it is desirable to avoid crossing of the freight trains with trains of the opposite direction in single-track sections, because of necessity of stop. the line benešov – české budějovice has not much passenger traffic, but three significant peaks. thus, double or triple traction for running through the whole line is worth consideration. 3.3. single-track lines busy national mainlines with mixed traffic: • domažlice – plzeň jižní předměstí, • brno – holubice – nezamyslice – přerov. mainlines with significant freight traffic: • české budějovice – horní dvořiště (– linz). middle-busy mainlines: • cheb – plzeň (there are double-track sections cheb – lipová u chebu and pňovany – plzeň only in ends of the line, so there is a character of a single-track line), • (furth im wald) – domažlice, • ústí nad orlicí – lichkov (– wroclaw). mainlines with little traffic: • české budějovice – české velenice (– gmünd nö). on single-track lines, longer overtaking tracks in almost every station are necessary, because of crossing of a freight train with any train which runs in the opposite direction. the alternative, which may be more efficient, is targeted partial doubling of the line. in the czech republic, most of passenger trains operate in (more or less) periodic timetable. if there are two systems of periodic passenger services with different section runtimes (and hence, with different crossing stations), there usually remains very little usable capacity left for freight trains. so, if there is a demand after a freight train paths approximately every 2 hours or more often, targeted partial doubling of the line is desirable. "targeted" means in the sections where trains from opposite direction cross each other. for the passenger trains in the czech republic, it usually corresponds to "zero symmetry" time, i.e. the minute 00. freight train paths should be, in the authors’ opinion, if possible, designed as periodic with identic symmetry time, and meeting of the freight 15 m. drabek, v. janos, z. michl acta polytechnica ctu proceedings trains of the opposite direction should be designed into the planned double-track section. if the freight traffic is irregular, at least crossing of a freight and a passenger train should be designed there. 3.4. sections and nodes with lack of long overtaking tracks there are listed lines, sections or nodes where additional long overtaking tracks are desirable: • cheb – plzeň, • furth im wald – domažlice – plzeň, • plzeň – české budějovice (a bypass line), • plzeň – beroun – praha (establishment of flank protection for long overtaking tracks can mostly solve the problem), • české budějovice – horní dvořiště (– linz) (a bypass line), • praha – české budějovice (a bypass line. after completed modernization, overtaking by fast trains is likely inevitable), • praha – ústí nad labem – děčín (this direction only), • děčín prostřední žleb – lysá nad labem, • kolín – havlíčkův brod – brno, • brno – přerov (modernization of the whole line is planned), • praha – olomouc (lack of long overtaking tracks outside large nodes), • přerov – ostrava (this direction only), • brno – břeclav (in the middle of the line and břeclav freight yard). 4. conclusion besides sections with high gradients, overtaking of a freight train is sometimes necessary because of fitting into spare time slot in a bottleneck area, or on a neighbouring line with a different timetable pattern. so, even on a line with sufficient capacity and train path homogeneity, a pair of long overtaking tracks is desirable for this reason. another reason is preparedness for traffic disruptions. in the case of higher gradients, such track should lay behind the peak of the line and before beginning of higher gradient uphill section, but neither in busy node nor in suburban or another bottleneck area. the discussion above has shown that for fulfilling eu requirements for enabling operation of 740 m long freight trains, it is not necessary to lengthen overtaking tracks in every station on double-track lines. but such tracks are desirable outside bottleneck areas and, if need be, behind a peak of the line. for the freight railway undertakings, the probably best capacity offer would be system of periodic catalogue freight train paths without unnecessary stops, as proposed by drábek [10], but with overtaking tracks long enough. this is the strategy that is recommended to sždc for long-term infrastructure improvement, as a conclusion of this paper. acknowledgements the authors would hereby like to propose a vote of thanks to sždc for enabling them to use railway station schemes [5] and especially to propose a vote of thanks to its employee dr. pavel krýže, who is the author of clearly arranged official maps of czech railway network [11]. the authors were glad to use the mentioned sources for the research presented in this paper. references [1] european commission. directive 2008/57/ec of the european parliament and of the council of 17 june, 2008 on the interoperability of the rail system within the community [online], 2011. http://eur-lex.europa.eu/legal-content/en/txt/ pdf/?uri=celex:02008l0057-20110322&from=en. [2] european commission. regulation (eu) no. 1315/2013 of the european parliament and of the council of 11 december 2013 on union guidelines for the development of the trans-european transport network and repealing decision no. 661/2010/eu [online], 2013. http://eur-lex.europa.eu/legal-content/en/all/ ?uri=uriserv:oj.l_.2013.348.01.0001.01.eng. [3] european commission. regulation (eu) no.913/2010 of the european parliament and of the council of 22 september 2010 concerning a european rail network for competitive freight [online]. european union, brussels, 2010. http://eur-lex.europa.eu/legal-content/en/ all/?uri=celex:32010r0913. [4] j. kušnír, j. ilík. railway in 2030, keynote speech, in: future of railway passenger transport in the czech republic, czech raildays [online], 2013. ostrava: czech ministry of transport, in czech, http://www.railvolution.net/czechraildays/2013/ seminare/konference-kusnir.pdf. [5] sždc: collection of official requisites for 2015/16 timetable, unpublished, praha: sždc, 2015. [6] r. brejcha, r. čech. operation of freight trains long up to 740 m. in: scientific-technical proceedings of czech railways. vol. 40/2015 [online], 2015. praha: czech railways, in czech, http://vtsb.cd.cz/vts/clanky/vts40/4008.pdf. [7] m. binko. railway infrastructure for freight transport, in czech raildays: adaptation of railway to both actual and future transport needs of passengers and goods [online], 2015. praha: m-presse plus s.r.o., in czech., http://www.railvolution.net/ czechraildays/2015/buletin2015.pdf. [8] m. binko. railway infrastructure for freight transport, keynote speech, in czech raildays conference [online], 2015. ostrava: sždc s.o., [2015-06-16], in czech, http://binko.webzdarma.cz/2015-6b.pdf. [9] czech technical standard tnž 34 2620 railway station and line signalling and interlocking systems, amended, effective since july 1. 2002, unpublished, olomouc: české dráhy, 2002. [10] m. drábek. periodic freight train paths in network, doctoral thesis. 2014. praha: ctu in prague, http://takt.fd.cvut.cz/cargo/drabek_thesis.pdf. [11] p. krýže. maps of railway network [online], 2015. praha: sždc, http://provoz.szdc.cz/portal/ viewarticle.aspx?oid=594598. 16 http://eur-lex.europa.eu/legal-content/en/txt/pdf/?uri=celex:02008l0057-20110322&from=en http://eur-lex.europa.eu/legal-content/en/txt/pdf/?uri=celex:02008l0057-20110322&from=en http://eur-lex.europa.eu/legal-content/en/all/?uri=uriserv:oj.l_.2013.348.01.0001.01.eng http://eur-lex.europa.eu/legal-content/en/all/?uri=uriserv:oj.l_.2013.348.01.0001.01.eng http://eur-lex.europa.eu/legal-content/en/all/?uri=celex:32010r0913 http://eur-lex.europa.eu/legal-content/en/all/?uri=celex:32010r0913 http://www.railvolution.net/czechraildays/2013/seminare/konference-kusnir.pdf http://www.railvolution.net/czechraildays/2013/seminare/konference-kusnir.pdf http://vtsb.cd.cz/vts/clanky/vts40/4008.pdf http://www.railvolution.net/czechraildays/2015/buletin2015.pdf http://www.railvolution.net/czechraildays/2015/buletin2015.pdf http://binko.webzdarma.cz/2015-6b.pdf http://takt.fd.cvut.cz/cargo/drabek_thesis.pdf http://provoz.szdc.cz/portal/viewarticle.aspx?oid=594598 http://provoz.szdc.cz/portal/viewarticle.aspx?oid=594598 acta polytechnica ctu proceedings 5:12–16, 2016 1 introduction 2 methods 2.1 rules for determination of available overtaking tracks for 740 m long trains 2.2 where to stop for overtaking – a soft decision process 3 results and discussion 3.1 double-track (or triple-track) lines 3.2 partially double-track lines (middle-busy mainlines) 3.3 single-track lines 3.4 sections and nodes with lack of long overtaking tracks 4 conclusion acknowledgements references 215 acta polytechnica ctu proceedings 1(1): 215–221, 2014 215 doi: 10.14311/app.2014.01.0215 highly magnetized accreting pulsars: are there accreting magnetars? pere blay1, pablo reig2, vı́ctor reglero1 1ipl, university of valencia, spain 2university of crete, greece corresponding author: pere.blay@uv.es abstract 2s 0114+650, gx 301-2, igr j16358-4726, x per, 4u 2206+54, sxp 1062, and 3a 1954+319 are thought to possess high magnetic fields. they have recently been named accreting magnetars, or highly magnetized accreting pulsars. in this work their properties are reviewed. within the context of their observational properties (mainly from integral data), and the recent models of accretion onto highly magnetized neutron stars, their similarities and differences are analyzed. the aim is to find a common framework to understand the evolution (in terms of past and present history) of these sources, and to establish the basis of a possible new kind of accreting sources. two of these sources, namely x per and 4u 2206+54, contain a massive main-sequence companion, while the rest are supergiant x-ray binaries or symbiotic systems. the variety of astrophysical scenarios represented by this set is wide, therefore the study of these systems is also important in order to establish commonalities between the different types of accreting x-ray pulsars and to study the possible evolutionary links between them. keywords: stars: binaries: close stars: magnetars stars: neutron x-rays: binaries. 1 introduction magnetars are neutron stars with bright persistent xray emission (on the order of lx ∼ 1033−36ergs−1) and spin periods on the order of a few seconds. they show a very drastic spin period evolution, with large spin period derivatives and a long-term spin down tendency. the fact that they are isolated objects, together with the difficulties in considering their rotational energy as the source of the observed bright x-ray emission, are key issues in order to postulate their highlymagnetized nature (with extreme magnetic fields, on the order or greater than ∼ 1014gauss). magnetars are usually grouped into two subclasses, namely anomalous x-ray pulsars (axps) and soft gamma-ray repeaters (sgrs). for a review of magnetars see, for example, panchapakesan (2003) and rea & esposito (2011). high mass x-ray binary systems (hmxrbs) contain a massive companion (≥10m�) and a neutron star or a black hole. the evolution of neutron stars in hmxrbs is affected directly by the wind plasma in which they are immersed (originating from the massive companion). the ambient plasma interacts with the magnetic field of the neutron star modifying its spin period evolution. neutron stars are born as rapid rotators. the ambient plasma removes rotational energy from the neutron star and transports it outwards and the neutron star suffers an spin-down process, which goes on until the ambient matter becomes radiative, cools rapidly, and accretion can take place. the time spent in this initial spin-down epoch will determine the final x-ray pulsar rotational period which will be an equilibrium period which resulting from the balance of two forces: the magnetic field pressure (which stops and drags the in-falling matter), and the inward pressure exerted by the ambient plasma. initial modelings (see, for example, illarionov and sunyaev, 1975) predicted that the slowest reachable pulse period of a neutron star in a binary system is on the order of ∼500 s. revisions of these initial models (see, for example, davies and pringle, 1981) relaxed this limit, allowing longer spin periods but only associated with higher neutron star magnetic fields. the possible presence of magnetars in some accreting hmxrbs, in which a slowly rotating neutron star has been found, has been proposed recently by a number of authors (see, for example, li & van den heuvel 1999, finger et al. 2010, reig et al. 2012, ikhsanov & beskrovnaya 2010, doroshenko et al. 2010, patel et al. 2007, popov & turolla 2012, bozzo et al. 2008). our goal is to compare the observational properties of these systems and determine which observational parameters are more suitable to derive the strength of the magnetic field. we also want to stress the difficulties in determining some of the properties of these systems when 215 http://dx.doi.org/10.14311/app.2014.01.0215 pere blay, pablo reig, vı́ctor reglero considering only the long-term analysis (i.e., comparing pulse periods from different epochs), or the temptation to identify spectral features as cyclotron resonant scattering features (crsf) and derive magnetic field values from their central energy. despite the small number of systems, the variety of astrophysical scenarios represented by this set is wide, from main sequence accreting binary systems (x per, sxp 1062 and 4u 2206+54) to supergiants (2s 0114+65 and gx 301-2), and even symbiotic systems (igr j16358-4726 and 3a 1954+319). in this sense, the study of these systems is important also to establish commonalities between the different types of hmxrbs and to study the possible evolutionary links between them. in this work we will review the observational properties of some of these systems and we will try to establish the scientific rational to develop a deeper analysis of their common properties and evolutionary status. due to the large inhomogeneity in the energy ranges and sets of data available in the literature, we have started an analysis of these systems by using integral/ibis/isgri1 data, setting up common energy ranges, spectral resolutions, and performing a homogeneous timing analysis. 2 the candidates a number of hmxrbs have already been proposed to contain highly magnetized neutron stars, namely 2s 0114+65 (li & van den heuvel 1999), 4u 2206+54 (reig et al. 2012, finger et al. 2010), x per (doroshenko et al. 2012), gx 301-2 (ikhsanov & finger 2012), igr j16358-476 (patel et al. 2007). there are also suggestions that sxp 1062 (fu & li 2012) and 3a 1954+319 (marcu et al. 2011) contain a highly magnetized neutron star. attempts have also been made to try to explain the properties of sfxts with the presence of a highly-magnetized neutron star with some degree of success (see bozzo et al. 2008 ) . a complete list of sources has been compiled in table 1. despite the fact that there are peculiarities in each of these systems which make difficult to group them as a kind (see, for example, reig et al. 2012, blay and reglero 2013), we can summarize their overall commonalities: • slowly rotating neutron stars (pspin slower than 500 s), with • high magnetic fields (larger than 1013 gauss) • most of them show a long-term spin-down tendency the neutron star properties in these systems are different from those of magnetars, for this reason the terms magnetar-like, or magnetar-descendant (making reference to a possible evolution from a neutron star born as a magnetar), have also been used to identify these sources (reig et al. 2012, blay and reglero 2013). table 1: list of sources. source pspin (s) porb (d) type 4u 2206+54 5560 19.2 ms donor x per 864 250 bexrb 2s 0114+650 9700 11.6 sxrb gx 301-2 680 41.498 sxrb igr j16358-4726 5880 −− symbiotic 3a 1954+319 19080 −− symbiotic sxp 1062 1062 ∼300 bexrb after 10 years of operations of the integral satellite, a very complete and homogeneous long-term database of observations of all these sources has been compiled. we have analysed all public observations in which 4u 2206+54, x per, 2s 0114+65, igr j16358476, and gx 301-2 were in the field of view of integral/ibis/isgri. the behavior of these source is similar in long time scales, with transient emission in the form of bursts or peaks and more or less continuous detection. they all show the typical behavior of wind accreting systems (although x per is a bex, it has been proposed that its behavior resembles that of wind-fed systems because of the large orbital distance, which will keep the neutron star away from the be disk, see doroshenko et al. 2012). the light curves, in every case, are complete enough to develop both short and long-term analysis of the spectral and timing properties of these systems. the magnetic properties of the neutron stars in these systems have been investigated/explained mainly in these terms: • spin period (neutron star evolutionary considerations) • spectral features (cyclotron resonant scattering features, crsf) 1integral (international gamma-ray laboratory) is an esa mission with contributions from usa and russia, with two high-energy imagers (ibis, working in the energy range 15 kev−10 mev , and jem-x sensitive in the 3−35 kev energy range) and one spectrometer (spi, sensitive in the range 20 kev−8 mev). ibis, in turn, has two detector layers, isgri, for the lower energy band,15 kev−1 mev, and picsit working in the 175 kev−20.4 mev energy range. a detailed description of the mission can be found in winkler et al. (2003) 216 highly magnetized accreting pulsars: are there accreting magnetars? with regard to the spin-period history, according to ikhsanov (2007), the maximum spin period reachable by a neutron star can be related to the intensity of its magnetic field and the quantity of accreted matter (i.e., mass accretion rate). the latter quantity is considered to be directly proportional to the x-ray luminosity of the source (see, for example, finger et al. 2010). the relationship between these quantities is: pmax = 15000µ 16/21 32 m −4/21 ns ( ṁ 1015gs−1 )−5/7s (1) where µ32 is the dipole magnetic moment (in units of 1032gcm−3) and mns the mass of the neutron star (in units 1.5m�), and ṁ is the mass accretion rate. a more recent calculation of quasi-spherical accretion by popov and turolla 2012, yields this relationship, directly in terms of the magnetic field intensity: b12 ∼ 8.1ṁ 1/3 16 v −11/3 300 ( p1000 porb,300 )11/2 (2) where b12 denotes the magnetic field in units of 1012 gauss, ṁ16 is the mass accretion rate in units of 1016gs−1, v300 is the wind velocity in units of 300kms−1, p1000 is the spin period in units of 1000s, and porb,300 is the orbital period of the system in units of 300d. on the other hand, a direct measurement of the magnetic field of the neutron star is possible when its high-energy spectrum shows crsfs. crsfs are due to the splitting of energy levels of the electron in the presence of a magnetic field. therefore the scattering of x-ray photons by these electrons is produced at quantized landau levels producing the absorption-like features seen in the x-ray spectra of many hmxrbs (see schönherr et al. 2007). the magnetic field strength can be calculated from the position of the crsf by the formula: ecrsf = 11.6 b12 (1 + z) (3) where z is the gravitational redshift. it should be noted that this redshift, and consequently the measured magnetic field, will depend on the height over the neutron star surface where the crsf is formed. in the next sections, we will review these methods when applied mainly to some of sources listed in table 1. 3 magnetic field determination via spin period evolution fig. 1 shows examples of spin period histories of 4u 2206+54, 2s 0114+650, igr 16358-4726, and x per. 53600 53700 53800 53900 54000 54100 54200 54300 mjd 5300 5400 5500 5600 5700 5800 p sp in ( s) 4u 2206+54 53000 53500 54000 54500 55000 55500 56000 mjd 2.00 2.25 2.50 2.75 3.00 p sp in (h ) 2s 0114+650 54000 54200 54400 54600 54800 55000 55200 55400 mjd 830 831 832 833 834 835 836 837 838 839 840 841 842 843 p sp in ( s) x per 53000 53500 54000 54500 55000 55500 56000 mjd 5780 5800 5820 5840 5860 5880 5900 p sp in ( s) igr j16358-4726 figure 1: spin period evolution of 4u 2206+54 (top panel), 2s 0114+650 (second panel from the top), x per (third panel from the top), and igr 16358-4726 (lower panel) as obtained from integral/ibis/isgri data in the 20-40 kev energy range. 217 pere blay, pablo reig, vı́ctor reglero 0.0001 0.001 0.01 0.1 1 f lu x (c ou nt s -1 k ev -1 ) 20 40 60 80 100 120 140 160 180200 energy (kev) -0.04 -0.02 0 0.02 0.04 re si du al s 4u 2206+54 0.01 0.1 1 f lu x (c ou nt s s1 ke v -1 ) 20 40 60 80 100 120 140 160 180200 energy (kev) -0.4 -0.2 0 0.2 0.4 re si du al s 2s 0114+650 0.0001 0.001 0.01 0.1 1 f lu x ( co un t s1 ke v 1 ) 20 30 40 50 60 70 80 90 100 200120 140 160 180 energy (kev) -0.02 -0.01 0 0.01 0.02 re si du al s x per figure 2: the 20-200 kev spectra of 4u 2206+54 (top panel), 2s 0114+650 (middle panel) and x per (bottom panel) extracted from integral/ibis/isgri data. the spectrum of 4u 2206+54 has been fitted with a bulk comptonization model with a χ2red ∼ 1.4, no evidence of a feature around ∼30 kev can be seen, neither in the spectrum nor in the residuals. for 2s 0114+650 a bulk comptonization model has been used to fit the data with a χ2red ∼ 1.2, again no evidence of a crsf at ∼30 kev is evident. for the case of x per, a comptonization model and two absorption features at 31.2 kev and 82.3 kev (indicated by arrows) have been used to fit the observed spectrum, with a χ2red ∼ 1.1. in the case of 4u 2206+54, frequency derivative values derived by finger et al. (2010) and reig et al. (2012) are 1.5 × 10−14hz s−1 and 1.7 × 10−14hz s−1 respectively. finger et al. 2010 included an older bepposax measurement. from these spin-down measurements, finger et al. (2010) estimated a magnetic field of b∼ 1014 g. by using the relationship of ikhsanov (2007) they also derived a magnetic field for 4u 2206+54 on the order of 1014 g. reig et al. (2012) calculated a magnetic field strength on the same order of magnitude by considering the relationship of popov and turolla (2012). the upper panel of fig. 1 shows the spin period measurements from integral/isgri data of 4u 2206+54 in the period 53600 − 54300 mjd. during this epoch (see blay and reglero 2013) the data is compatible with no changes in the spin period of the system, as the pulse period measurements show large errors. the same thing happens for x per and 2s 0114+650 (second and third panels from the top, respectively), in which large errors in the spin period determination make the measurements compatible with no pulse period change. therefore we cannot report on magnetic field measurements from spin period changes for these three systems. the luminosity of 2s 0114+65 in the epoch shown in fig. 1 is 4 × 1037ergs−1 (estimated from the model fitted to the spectrum shown in fig. 2), considering a distance to the source of 7.2 kpc (hall et al. 2000). therefore, the ikhsanov (2007) relationship yields a magnetic moment on the order of 1033 g cm−3, which yields a magnetic field strength of b ∼ 1015 g. the relationship from popov and turolla (2012), however, results in an unreasonable number when applied to this source. doroshenko et al. (2012) argue that, for the case of x per, according to the torques applied to the neutron star magnetosphere in the case of wind accretion, and taking into account the theoretical approaches of doroshenko et al. (2010), davidson and ostriker (1973), davies et al. (1979), and bisnovatyi-kogan (1991), the magnetic field in x per is expected to be on the order of b∼ 1014. we show in the lower panel of fig. 1 the spin period measurements made with integral/isgri data. for this source we can estimate the luminosity to be lx ∼ 1.09 × 1035ergs−1 in this epoch (estimated from the model fitted to the spectrum shown in fig. 2 ). as the luminosity of the source is closely related to the accretion rate, we can estimate the field strength of x per by using the relationship of popov and turolla (2012), yielding b∼ 1012 g, which falls within the range of normal neutron star magnetic fields. the relationship from ikhsanov (2007) results in a magnetic moment of µ ∼ 1030 g cm−3, which implies a magnetic field on the order of b ∼ 1012 g. we find, therefore, a discrepancy with the results of doroshenko et al. (2012). we cannot reproduce their calculations, however, because of the large errors in the spin period determination, which prevent the determination of a reliable frequency derivative, needed for the determination of the spin torques. for igr j16358-4726 we see in the lower panel of 218 highly magnetized accreting pulsars: are there accreting magnetars? fig. 1 that the source experiences large pulse period changes. it shows a spin-down between mjd∼53250 and mjd∼54000 of ν̇ ∼ 2.4 × 10−14 hz s−1. according to the equation (9) from finger et al. (2010), which relates the magnetic moment of the neutron star with the measured spin-down, we find in igr j16358-4726 a minimum magnetic moment of µ ∼ 2 × 1032 g cm−3, which implies a magnetic field on the order of or larger than b∼ 1014 g. these calculations rely on a good measurement of the luminosity (or mass accretion rate) or of the changes in the spin period. the major uncertainty comes from the determination of the luminosity of the source, calculated from the source flux in a given energy range and the estimated distance to the source, which is usually subject to large uncertainties. with respect to the use of spin period changes, the estimates always rely on long-term or average behavior. on the one hand this is due to the poor knowledge of the short-term or detailed pulse period evolution in many cases, and, on the other hand, to the lack of a detailed and complete theoretical description of the wind accretion scenario. the efforts of postnov et al. (2013), however, try to fill this lack of good theoretical approximations. we will comment on that in the last section. 4 magnetic field determination via cyclotron resonant scatering features a more direct measurement of the magnetic field strength of the neutron star in these systems can be obtained by determining the position of the fundamental crsf seen in their x-ray spectra. crsf have been reported in gx 301-02 (la barbera et al. 2005, makishima and mihara 1992), 2s 0114+65 (wang 2010, bonning and falanga 2005), 4u 2206+54 (blay et al. 2005, , wang 2009, torrejón et al. 2004, masseti et al. 2004), and x per (this work, see fig. 2, doroshenko et al. 2010), see table 2. table 2 summarizes the crsf reported in these systems and the magnetic field strength measurement derived. table 2: crsfs and b determination. see references in the text. source ecrsf (kev) b (10 12 gauss) 4u 2206+54 30 2.6 x per 31.2 2.7 2s 0114+650 22 1.9 gx 301-2 37 3.2 5 2s 0114+650 22(c) 1.9 we see in all cases that the magnetic field strength derived by this method is on the order of those found in other neutron stars in hmxrbs, i.e., 100 to 1000 times lower than those of the magnetars. we find, therefore, a discrepancy in magnetic field strength determination with respect to the methods shown in the previous section. one could think that the magnetic fields derived from spin period evolution are not properly calculated, in view of the more direct determination from crsf measurements. however the location of crsf is not free of ambiguities. in the case of 4u 2206+54 (see top panel of fig. 2), for example, this feature has been observed only marginally, and only at one epoch (see blay et al. 2005 and references there in). it has not been observed again (see blay and reglero 2013). therefore, although it has been assumed that the absorption observed was a crsf, this may not be the case. a similar situation applies to 2s 0114+65 (see middle panel of fig. 2), in which the detection of the possible crsf has only been marginal (see wang 2010) and not confirmed again. in the case of x per, doroshenko et al. (2010) show how instead of a crsf fundamental and its first harmonic (see bottom panel of fig. 2), the x per spectra can be very well described by a combination of two comptonization models (comptt+comptt, in xspec notation) demonstrating the possibility that the features observed may not be crsfs. therefore, at least in some cases, the possibility of finding magnetar-like magnetic fields in these systems is not ruled out by the reported observations of crsfs. 5 discussion and conclusion we have seen that in most cases the determination of the magnetic field strength of the neutron star is ambiguous, or even contradictory, when the result from different approaches are compared. we want to emphasize how different approaches may seem to result in different magnetic field strength estimates. an homogeneous and coherent approach is needed in order to determine if there are links between these systems (in terms of their spin period and binary evolution) and to support or discard the different theories trying to explain the puzzling presence of long spin periods in hmxrbs. there have been recent theoretical efforts in order to provide better modeling of the wind-accretion scenario. ikshanov and beskrovnaya (2013), for example, report on a likely explanation of the large spin-down shown in 4u 2206+54 by taking into account a higher than usual magnetic field of the companion star and, consequently, that the accreting material can be magnetized. the possibility of an optical companion with a large magnetic field had been already proposed by blay and reglero (2011). 219 pere blay, pablo reig, vı́ctor reglero the efforts of postnov et al. (2013) in modeling more accurately the wind accretion mechanism in hmxrbs also result in magnetic field estimations for 4u 2206+54 and sxp 1062 on the order of typical magnetic fields found in accreting x-ray pulsars (∼ 1012g). they explain the behavior of sources like gx 301-2 with the hypothesis of being older systems which have already reached their equilibrium period. together with an improvement in the modeling of wind accretion in hmxrbs, a better observational approach is needed in order to understand the behavior of these sources. multiwavelength campaigns (ir, optical, uv, x-rays) on these objects are needed (as simultaneous as possible) in order to determine the various parameters involved: obtain more accurate spectral classification and orbital solutions, measure accurate mass loss rates and wind velocity laws, etc. acknowledgement p. blay acknowledges funding from the spanish ministerio de economia y competitividad through project aya-2011-29936-c05. references [1] bisnovatyi-kogan, g. s. 1991, a&a, 245, 528 [2] blay p., ribó m., negueruela i., torrejn j. m., reig p., camero a., mirabel, i.f., & reglero v. 2005, a&a 438, 936 [3] blay p. and reglero v. 2011, bsrsl, 80, 634 [4] blay p. and reglero v. 2013, pos(integral 2012)012 [5] bonning e. w., falanga m. 2005, a&a, 436, l31 [6] bozzo e., falanga m., stella l. 2008, apj, 683, 1031 doi:10.1086/589990 [7] davidson, k., and ostriker, j. p. 1973, apj, 179, 585 [8] davies, r. e., fabian, a. c., and pringle, j. e. 1979, mnras, 186, 779 [9] davies r.e., and pringle j.e. 1981, mnras, 196, 209 [10] doroshenko v., santangelo a., suleimanov v., kreykenbohm i., staubert r., ferrigno c., klochkov d. 2010, a&a 515, a10 [11] doroshenko v., santangelo a., kreykenbohm i., and doroshenko r. 2012, a&a 540, l1 [12] finger m.h., ikhsanov n.r., wilson-hodge c.a., & patel s.k. 2010., apj, 709, 1249 doi:10.1088/0004-637x/709/2/1249 [13] fu l., li x.d. 2012, apj, 757,171 doi:10.1088/0004-637x/757/2/171 [14] fürst f., marcu d.m., pottschmidt k., grinberg v., wilms j., and cadolle bel m. 2011, pos(integrall 2010)017 [15] hall t. a., finley j. p., corbet r. h. d., thomas r. c., 2000, apj, 536, 450 doi:10.1086/308924 [16] harbel, f., sturm, r., & filipović, m. d. et al. 2012, a&a, 537, l1 [17] ikhsanov, n. r. 2007, mnras, 375, 698 doi:10.1111/j.1365-2966.2006.11331.x [18] ikhsanov n.r., finger m.h. 2012, apj, 753,1 doi:10.1088/0004-637x/753/1/1 [19] ikhsanov n.r., beskrovnaya n.g. 2010, astrophysics, 53-2, 237 [20] ikhsanov n. r. and beskrovnaya n. g.2013, arep, 57, 287 [21] illarionov a.f., and sunyaev r.a. 1975, a&a, 39,185 [22] koh d.t., bildsten l., chakrabarty d., et al. 1997, apj, 479, 933 [23] li x. d., van den heuvel e. p. j., 1999, apj, 513, l45 [24] la barbera, a., segreto, a., santangelo, a., kreykenbohm, i., & orlandini, m. 2005, a&a, 438, 617 [25] makishima k., mihara t. 1992, in magnetic fields of neutron stars, ed. t. tanaka & k. koyama (tokyo: universal academy press), 23 [26] marcu d. m., fürst f., pottschmidt k., et al. 2011, apj, 742, l11 doi:10.1088/2041-8205/742/1/l11 [27] masetti n., dal fiume d., amati l., del sordo s., frontera f., orlandini m., and palazzi e. 2004, a&a, 423, 311 [28] panchapakesan n. 2003, bull. astr. soc. india, 31, 19 [29] patel s.k., zurita j., del santo m. et al. 2007, apj, 657, 994 doi:10.1086/510374 [30] popov s.b., turolla r. 2012, mnras, 421, l127 doi:10.1111/j.1745-3933.2012.01220.x 220 http://dx.doi.org/10.1086/589990 http://dx.doi.org/10.1088/0004-637x/709/2/1249 http://dx.doi.org/10.1088/0004-637x/757/2/171 http://dx.doi.org/10.1086/308924 http://dx.doi.org/10.1111/j.1365-2966.2006.11331.x http://dx.doi.org/10.1088/0004-637x/753/1/1 http://dx.doi.org/10.1088/2041-8205/742/1/l11 http://dx.doi.org/10.1086/510374 http://dx.doi.org/10.1111/j.1745-3933.2012.01220.x highly magnetized accreting pulsars: are there accreting magnetars? [31] postnov k.a., shakura n.i., kochetkova a.yu., and hjalmarsdotter l., 2013, arxiv:1307.3032 [32] rea n., esposito p. 2011, in high-energy emission from pulsars and their systems, ed. d. f. torres & n. rea, astrophysics and space science proceedings (springer berl in heidelberg), 247 [33] reig p., torrejón j.m., & blay p. 2012, mnras, 425, 595 doi:10.1111/j.1365-2966.2012.21509.x [34] schönherr g., wilms j., kretschmar p., kreykenbohm i., santangelo a., rothschild r.e., coburn w., staubert r. 2007, a&a 472, 353 [35] torrejn j. m., kreykenbohm i., orr a., titarchuk l., and negueruela i. 2004, a&a, 423, 301 [36] wang, w. 2009, mnras, 398, 1428 doi:10.1111/j.1365-2966.2009.15200.x [37] wang w. 2010, a&a 516, a15 [38] winkler c., courvoisier t.j.l., di cocco g., et al. 2003, a&a, 411, l1 221 http://dx.doi.org/10.1111/j.1365-2966.2012.21509.x http://dx.doi.org/10.1111/j.1365-2966.2009.15200.x introduction the candidates magnetic field determination via spin period evolution magnetic field determination via cyclotron resonant scatering features discussion and conclusion 41 acta polytechnica ctu proceedings 2(1): 41–45, 2015 41 doi: 10.14311/app.2015.02.0041 cvs around the minimum orbital period s. zharikov1, g. tovmassian1 1observatorio astronomico nacional, instituto de astronomia, universidad nacional autonoma de mexico, ensenada, bc, mexico, 22860 corresponding author: zhar@astrosen.unam.mx abstract we discussed features of cataclysmic variables at the period minimum. in general, most of them must be wz sge-type objects. main characteristics of the prototype star (wz sge) are discussed. a part of wz sge-type objects has evolved past the period limit and formed the bounce back systems. we also explore conditions and structure of accretion disks in such systems. we show that the accretion disk in a system with extreme mass ratio grows in size reaching a 2:1 resonance radius and are relatively cool. they also become largely optically thin in the continuum, contributing to the total flux less than the stellar components of the system. in contrast, the viscosity and the temperature in spiral arms formed at the outer edge of the disk are higher and their contribution in continuum plays an increasingly important role. we model such disks and generate light curves which successfully simulate the observed double-humped light curves in the quiescence. keywords: cataclysmic variables dwarf novae period minimum. 1 introduction a widely accepted evolutionary theory of cataclysmic variables (cvs), as presented in kolb & baraffe (1999, and references therein), predicts a significant accumulation of cv systems around the orbital period minimum (paczynski 1981). it also envisions that ∼70% of the current cv’s population has evolved past the orbital period minimum and formed so-called bounce-back systems. figure 1 illustrates the current concept of cv evolution at the orbital period turn-around point on the mass-transfer rate and mass-ratio to orbital period diagrams. cataclysmic variables with orbital periods close to the 80 min orbital period minimum that undergo infrequent (years to decades) super-outbursts are called wz sge-type stars. objects with short periods that have not been observed in outburst or super-outburst but have spectral characteristics similar to wz sge are listed also as wz sge-type candidates. those systems are dominated at the period minimum. the bounce-back systems are cataclysmic variables evolved beyond the minimum period limit, which is reached when the secondary star becomes of a substellar mass (brown dwarf) and partially degenerate. bounce-back systems are expected to float within the 80 100 min orbital period range. bounceback systems are spectroscopically similar to these, but not every wz sge-type object has necessarily passed through the turning point. figure 1 also displays the expected position of the bounce-back systems. in this paper we discuss observational characteristics of cataclysmic variables at the period minimum and explore conditions and structure of its accretion disk. figure 1: schematic distribution of cv-types on the plot of the mass ratio and mass transfer rate vs. the orbital period. 2 wz sge-type stars wz sge is the prototype star of the class of short-period cataclysmic variables (named wz sge-type stars). it is a high inclination (∼77o; steeghs et al. 2007) cataclysmic variable with a late m-type dwarf secondary star (0.078m� < m2 < 0.13m�) orbiting at 81.6 min the fastest-spinning (∼ 28 sec, robinson et al 1978) 41 http://dx.doi.org/10.14311/app.2015.02.0041 s. zharikov, g. tovmassian white dwarf (0.88m� < m1 < 1.53m�). the distance to wz sge is only 43.5 pc. for most of its life it is in quiescence with a v ∼ 15, corresponding to mv ∼ 12. below the main characteristics of wz sge are summarised: • a short orbital period of 81.6 min, close to the predicted period minimum of cvs ∼ 77min with a main sequence secondary. • a spectrum in quiescence shows strong double-peaked balmer emission lines from the accretion disk surrounded by broad absorptions, formed by the primary white dwarf (see for an example howell et al. (2008)). • an infrequent ∼20-30yr and largeamplitude (∼ 8mag; 1913, 1946, 1978, 2001 yy.) super-outbursts succeeded by echo outbursts, absence of normal (as in su umatype stars) outbursts. in order to avoid an earlier occurrence of normal outburst in wz sge-type systems, it can be accepted as an extreme low viscosity parameter a ∼0.010.001 in accretion disks (smak (1993), osaki (1994)). • the light curve during a super-outburst shows long-lasting super-humps (patterson, et al. 2002). • the optical light curve are double-humped sometimes during super-outburst and in quiescence (patterson, et al. 1998; patterson, et al. 2002). • there is evidence of forming spiral arms in the disk during super-outburst (baba et al. 2002, howell et al. 2003). • in quiescence the accretion disk is asymmetric, and the bright spot region is shown to be extended along the mass transfer stream (skidmore et al. 2000, mason et al. 2000). • the outer layers of the accretion disk must be of low density and low temperature ∼3000k (howell et al. 2004). • a cavity most likely formed in the inner part of the disk during quiescence implying an annulus-shaped accretion disk (kuulkers et al. 2011). • the outer radius of the disk is about a 3:1 resonance radius (rdisk ≤ r3:1) in quiescence and it can reach to 2:1 resonance radius (rdisk ≤ r2:1) during a super-outburst. there are about ∼100 objects proposed or confirmed as wz sge-type systems. most of them were suggested based on features of spectra or the amplitude of the super-outbursts. figure 2 summarise current conception about wz sge-type systems which do not yet pass the period minimum. figure 2: the current conception about wz sgetype systems which do not yet pass yet the period minimum. 3 bounce-back systems after reaching the period minimum the cvs should evolve back toward longer periods and form a so-called bounce-back or post-period minimum systems. the increasing of orbital periods of bounce-back systems is accompanied by a decline in mass transfer rate about an order of magnitude according to the existing models of cvs evolution (kolb & barrafe 1999, sirotkin & kim 2010). the list of the bounce back candidates are given in table 2 and figure 3 (zharikov et al. 2013). the optical spectra in quiescence of proposed candidates are similar to wz sge. all those systems include a massive and relatively cool (∼ 12000k) white dwarf primary and a m2/m1 mass ratio less than 0.075. such low mass ratio implies a jupitersize secondary (a late type m-dwarf or a brown dwarf) and with a 2:1 resonance radius is within of the roche lobe of the primary. also, similar to wz sge, the bounce back candidates do not show normal outbursts which implies an extremely low α ∼ 0.01 − 0.001 viscosity parameter (smak (1993), osaki (1994)). this together with the low ∼ 10−11m�/year mass transfer rate, allows the accretion disk to expand in bounce back systems up to a 2:1 resonance radius and form a two spiral wave structure in those systems (lin & papaloizou, 1979). the large size of the disk and domination of white dwarf radiation in the optical range and the secondarys radiation in the jhk bands in sdss0804 and sdss1238 bounce back candidates (zharikov et al 2013, aviles et al 2010) lead to the conclusion that the ”standard accretion disk model” (frank et al. 2002) does not apply to bounce-back systems. 42 cvs around the minimum orbital period table 1: parameters of wz sge and of bounce-back candidates. nn/object porb v q m1 m2 t w d eff i lc 1 (days) (mag) (m�) (m�) (k) ( o) wz sge 0.0567 ∼ 15 0.092 0.85 0.078 13500 77 +sq 1. gw lib∗ 0.0533 19.1 0.060 0.84 0.05 11 2. v455 and∗ 0.0563 16.5 0.060s >m9 11500 83 +q 3. al com∗ 0.0567 19.1 0.060 16300 +q 4. sdss1035 0.057 18.7 0.055e 0.94 0.05 10100 83 5. sdss1238 0.056 17.8 0.05 ∼ 1.0 0.05 12000 ∼70 +q 6. sdss0804∗ 0.059 17.8 0.05s ∼ 0.9 0.045 13000 ∼ 70 +q 7. eg cnc∗ 0.060 18.8 0.035s 12300 +s 8. rx1050-14 0.062 17.6 <0.055v 13000 <65 9. gd552 0.0713 16.6 <0.052v <0.08 10900 <60 10. re1255 0.083 19.0 <0.064v >0.9 <0.08 12000 < 5 1 light curve (lc) features: ”+” lc shows a double-hump during the orbital period; ”s” during super-outburst; ”q” during quiescence; ”-” absent of double-humps in lc. ∗ objects which demonstrated wz sge-type superoutburst. figure 3: the systems with known mass ratio vs. orbital period are plotted. the left axes is the mass ratio and estimated mass transfer rates are given on the right axes. the bounce-back candidates are enclosed in a dash-dotted box. cannizzo & wheeler (1984) studied the vertical structure of a steady-state, α-model thin-accretion disk for an accreting object of 1 m�. they found that, for low accretion rates, the disk structure is optically thin and can be doublevalued with high(∼5000 k) and low(∼2000 k) temperature branches. for α > 0.1 a warm solution is possible in the inner region of the accretion disk, but disk annuli at larger radii will be in a cold state with t < 2000 k. only the low-temperature solution exists for α ≈ 0.1. as α decreases with temperature, the tendency to develop cold solutions in quiescence is enhanced. therefore, accretion disks in bounce back systems are most probably cool (∼2500 k). the optical light curve of the high inclination bounce back candidates shows permanently a doublehumped light curve (marked by ”+q” in table 2). the spiral arm structures were found from doppler tomography mapping in quiescence in two well studied example of bounce back candidates sdss1238 (aviles et al. 2010) and sdss0804 (zharikov et al. 2013). also important, that, because the mass transfer rate significantly decrease after the period minimum the size of the magnetosphere of the primary with faint magnetic field will increase with the decreasing of ṁ and a relatively faint (≤ 1mgauss) magnetic field is enough to form a cavity in the inner part of the accretion disk (zharikov et al 2013). taking into account all these features a geometrical model of bounce back system was constructed to explain the observed double-humped light curve in quiescence. the model takes into account the positions of the bright structures in the doppler maps, the large size of the accretion disk, and the description of the spiral density waves in hachisu et al. (2004). figure 4 presents the geometry used in the model (central panel) and grayscale images present the height of the disk (left) and the temperature distribution (right) of the accretion disk. 43 s. zharikov, g. tovmassian figure 4: from left to right: vertical thickness of the accretion disk, model configuration used to calculate the light curves of bounce-back systems and the temperature distribution of the model. figure 5: left: examples of double-humped light curve; right: examples of the light curve generated by the model. we calculated a variety of models using the typical average parameters of bounce-back systems presented in table 1 and the double-hump-shape light curve (fig. 5, (left panels) is easily reproduced by such models, two examples of which are shown in fig. 5, (right panels). 4 conclusions the galactic age is high enough that ∼ 70% of cvs will have reached the cvs orbital period minimum. in general, most of them must be wz sge-type objects. a part of wz sge-type objects have already passed the period limit and formed the bounce back systems. the structure of the disk must be changed when the system pass through the minimum. before the minimum, mass transfer rate is higher and the accretion disk is hotter and smaller. after the period limit it is larger and reaches a 2:1 resonance radius. the disk is cool (∼ 2500k) and has a cavity in the inner part and presents permanently existing spiral arms. the spirals should be denser and hotter that the rest of the disk. based on these features we propose the double-humped light curve in quiescence state of a system together with orbital period longer than the period limits and the mass ratio of q < 0.08 as the main manifestation of a wz-sge system which already passed through the minimum orbital period. when accretion has ended the cvs can form very close short period wd+bd systems, two examples of which are found recently (wd0137-349, burleigh et al. (2006) and nltt5306, steele et al. (2013)). acknowledgement we acknowledge papiit grants in-109209/in-103912 and conacyt grants 34521-e; 151858. references [1] aviles, a., et al. 2010, apj, 711, 389 [2] baba, h. et al. 2002, pasj, 54, l7 44 cvs around the minimum orbital period [3] burleigh, m. r., et al. 2006, mnras, 373, l55 doi:10.1111/j.1745-3933.2006.00242.x [4] cannizzo, j. k., & wheeler, j. c. 1984, apjs, 55, 367 [5] frank, j., et al. 2002, accretion power in astrophysics, cambridge university press doi:10.1017/cbo9781139164245 [6] hachisu, i. et al. 2004, apjl, 606, l139 doi:10.1086/421295 [7] howell, s. b., et al. 2003, a&a, 399, 219 [8] howell, s. b., et al. 2004, apjl, 602, l49 doi:10.1086/382481 [9] howell, s. b., et al. 2008, apj, 685, 418 doi:10.1086/590491 [10] kolb, u., & baraffe, i. 1999, mnras, 309, 1034 [11] kuulkers, e., et al. 2011, a&a, 528, a152 [12] lin, d. & papaloizou, j., 1979, mnras, 186, 799 doi:10.1093/mnras/186.4.799 [13] osaki, y. 1995, pasj, 47, 47 [14] mason, e., et al. 2000, mnras, 318, 440 [15] paczynski, b. 1981, acta astron., 31, 1 [16] patterson, j., et al 1998, pasp, 110, 403 [17] patterson, j., et al. 2002, pasp, 114, 721 [18] robinson, e. l., et al. 1978, apj, 219, 168 [19] sirotkin, f. & kim, w. 2010, apj, 721, 1356 doi:10.1088/0004-637x/721/2/1356 [20] skidmore, w., et al. 2000, mnras, 318, 429 [21] smak, j. 1993, acta astronomica, 43, 101 [22] steeghs, d., et al. 2007, apj, 667, 442 doi:10.1086/520702 [23] steele, p. r., et al. 2013, mnras, 429, 3492 doi:10.1093/mnras/sts620 [24] zharikov, s., et al. 2013, a&a, 549, a77 discussion raimondo baptista: the opening angle of the spiral structure in an accretion disk depends on the ratio of the local sound speed to a keplerian speed. the disk region containing the spirals must be quite hot in order for the spiral will be wide open as in your model. is it physically consistent to get wide spiral in the disk with such low temperature such as ∼ 2000 k. sergey zharikov: our model is a geometrical model only. based on proposed geometrical shape and size of the disk in bounce back systems we can explain observed double-humped light curves in quiescence. formation such spiral arm structure and physical conditions in such cool disk is not clear yet. the first approach to describe such disks was presented in paper of cannizzo, & wheeler, (1984). 45 http://dx.doi.org/10.1111/j.1745-3933.2006.00242.x http://dx.doi.org/10.1017/cbo9781139164245 http://dx.doi.org/10.1086/421295 http://dx.doi.org/10.1086/382481 http://dx.doi.org/10.1086/590491 http://dx.doi.org/10.1093/mnras/186.4.799 http://dx.doi.org/10.1088/0004-637x/721/2/1356 http://dx.doi.org/10.1086/520702 http://dx.doi.org/10.1093/mnras/sts620 introduction wz sge-type stars bounce-back systems conclusions acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0038 acta polytechnica ctu proceedings 12:38–41, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app analysis of stored data help to propose and generate new tracks václav jirkovský czech technical university in prague, faculty of transportation sciences, department of vehicle technology. konviktská 20, 110 00 prague correspondence: vaclav.jirkovsky@volny.cz abstract. during experiments on vehicle simulators a large amount of data is stored. on these data, it is possible to trace some similarities in the behavior of drivers in certain areas or when performing the same task. we can assume that if the driver performs a certain type of experiment, his behavior exhibits certain traits. these elements of common behavior can be used to create virtual track for experiments. which elements and how they can be used is described in this article. the algorithm for automatic creation of virtual track based on type of experiment is provided. it will help us to define the purpose of measurement and the track could be generated automatically. keywords: vehicle simulator, measurement, virtual track. 1. introduction conducting of experiments is one of the most important things related to the operation of driving simulators at the faculty of transportation sciences. there are many types of them. they are focused on investigating the driver behavior in fatique[1–3], testing systems for predicting microsleeps[4], humanmachine interaction[5], interaction with surrounding traffic[6, 7] and many more. during their performing the data are collected, which are mostly of a technical nature, and on the other hand, they give us information about the state of the monitored driver. all information about a vehicle, terrain and environment belong the first category. the data collected on a driver's body such as heart beat, eeg, eye view direction, blink rate etc. belong to the second category[8]. after the data evaluation we are able to describe a driver's behavior in particular states of experiment and determine a crisis situations or areas. hierarchical structure of measured data is shown on fig.1 figure 1. hierarchical structure of measured data 2. measured data analysis if the amount of the measured data is sufficiently large, we are able to detect with help of mathematical tools how the drivers react to a given event, incentive or transport situation. we can find out if their reactions are different or identical. it is possible, based on the analysis and using of appropriate tools, to predict the behavior of drivers on virtual tracks if these tracks are only modeled or created without performing of experiments with data collection[9]. on the basis of structure and character of the track, it is possible to determine the place and the moment when the driver will have to solve the crisis situation. but the opposite procedure will be applied. it means to design and construct a virtual track according to the requirements of the type of crisis. on thus obtained results we can create new scenarios with a specific focus. for example the results of the analysis shows that while operating the radio in rugged terrain, a large percentage of drivers do not hold the vehicle on an ideal path and run off the road. it's evident that the equally segmented track is suitable for testing similar devices. based on these derived addiction it is possible to design and construct virtual tracks with greater efficiency. it will also allow it to specialize in one specific purpose. the figure no.2 shows the experimental track. the yellow marks show the places where the value of deviation from the ideal track overlaps 1.5 m. this value indicates that the car doesn't stay in the correct road lane and crosses the middle line to the opposite road lane. this can cause a serious traffic situation with possibility of frontal collision. the chart shows that the deviation from the ideal track is caused by splitting the focus of the driver between the driving and the device manipulation. for more detailed evaluation we can divide the tasks into easy and difficult. an easy task can be done by one 38 http://dx.doi.org/10.14311/app.2017.12.0038 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 analysis of stored data help to propose and generate new tracks figure 2. experimental track or two clicks (touch). difficult task needs more time to finish. 3. identifying common elements of behavior during data analysis i've found out that the common behavior patterns could be observed for most types of experiments. with a very high probability we can assume that if the driver performs the measurements under the same conditions as the previous driver, his behavior will show the same elements. by examining the experiments i received the following connection: • assistive devices ragged track • research on driver fatigue monotonous, minimally curved track with minimal distractions • research driver's responsiveness to sound stimuli monotonous track • research driver's responsiveness to obstacles in riding complex traffic situations, poor visibility • driving precision special testing polygon (slalom, braking precision, ride along a line etc.) • overtaking research track with poor clarity (bend, curl, heavy traffic) individual connections determine the characteristics. with their help, i set up a knowledge base. that gives us an idea of which features driver's behavior occur in different types of experiments. that will tell us what behavior we expect on a particular experiment, but in the opposite case it will help us, when examining on the basis of activity to determine the shape of table 1. experimental characteristics the track. characteristics of each experiment can be found in the following table. experiment properties c om pl ex tr ac k r ec om m en de d ve lo ci ty su rr ou nd in g tr affi c li gh t tr affi c h ea vy tr affi c g oo d vi si bi lit y c or ru ga te d tr ac k o bj ec ts reaction speed on sound effects no no yes no no reaction speed on road obstacles yes yes no yes no yes the accuracy of the ride yes no no no driver's fatigue no 90130 km/h no yes no no assistance devices handling yes 50km/h no overtaking yes yes no yes no yes 4. proposal track by type of experiment a knowledge base is essential for the design of the track, which i described in the last chapter. to create the track i proposed algorithm, which is described below. the entire process can be divided into several steps: • the user determines variables, which aims to examine • distinguishing features are loaded from knowledge base • the track is automatically proposed based on those distinguishing features • the result is a polyline characterizing the road axis in space first the user enters input data. it is primarily a type of experiment. the next steps of propose are fully automatic and the user cannot affect them. based on input values the distinguishing features are obtained from the knowledge base. it means maximal velocity, curvature, longitudinal profile etc. these go to input of an algorithm as a specific values (arcs with radius from 100m to 350m, narrow lines of length from 200m to 1000m etc.). based on the input data the algorithm proposes the whole track. geometric properties of individual building elements comply with standard csn 73 6101. the result is a polyline that determines the 39 václav jirkovský acta polytechnica ctu proceedings points of the track. it can be subsequently imported into other software where it's finalized[10]. 5. the algorithm for automatic generation of experimental tracks the automatic generation of a track is as old as the game industry itself. the limitation of existing platforms did not allow the distribution of a large predefined content. ad-hoc algorithmic procedures were widely used to generate game content on-the-fly. it was called procedural content generation. now, when the distribution of games is not limited by memory space, the automatic generation of the content is still used mainly to reduce design costs. the principle of automatic generation of the track is based on evolutionary computation to evolve track for a simple two-dimensional car racing game[11]. the basic idea is to connect individual segments (lines, arcs) so that they don't cross the other parts. and they gradually point to beginning. the connection of the parts is strictly defined. for example between each line and arc must be placed a spiral. the segment properties are based on the rules mentioned above. additional parameters slope can be added to extend into 3d space to make virtual track more realistic[12]. 6. pseudo-code as described in section four, the type of track is dependent on the type of experiment. it means each track consists of segments with a specific property. the track properties are set at the beginning of the algorithm. in the next step the random segment is randomly selected from the segment list. the list consists of narrow lines of different length and arcs of different length and radius. the end point of the previous segment is the start point of a next segment. the algorithm takes into account the crossing of the segments. if the actual segment crosses the track it is not placed and other segment is chosen. in some cases the algorithm doesn't offer the closed track, therefore, it must be done manually. to reduce the occurrence of these cases, the number of right-handed arcs is bigger than the number of left=handed arcs. once the track has been built, it needs to be smoothen. in this step the spirals are placed between a line and an arc to provide a fluent transition. (1.) var property = settrackproperty() (2.) track=null (3.) object=choose random object from list (4.) object.properties=random(property) (5.) end_point=object.endpoint() (6.) start_point=object.startpoint() (7.) track.insert(object) (8.) do while (start_point != point) (9.) object2=choose random object from list (10.) object2.properties=random(property) (11.) object2.start_point=end_point (12.) point=object2.endpoint() (13.) if(collision(track, object2) == false and suitableposition(object2, start_point) == true) (14.) track.insert(object2) (15.) calculate parts connection (16.) if(distance(start_point, point) < minimal_distance) (17.) break (18.) else (19.) point=end_point (20.) loop (21.) connect point and start_point (22.) track.smooth() 7. conclusions the measurements on a vehicle simulator and data analysis confirmed that the use of assistance and media devices can directly cause a dangerous behavior of the driver. those devices should be designed to reduce driver's attention and thereby minimize dangerous situations on the road. thorough tests could help us to determine which devices are suitable for using in a car and which are not. the analysis also showed that there are connections between driver's behavior and track's properties and conditions of the measurement. the connections define rules that can be used during creating of the new tracks. in the future we will be able to define purpose of measurement and the track will be generated automatically. the described algorithm for creating a track is one of many possible. it's easy to understand and fulfills the purpose. a number of track were generated during testing. many of the tracks were useful and usable as a basis of a new created track. their use is one of the steps to improve the quality of experiments on vehicle simulators. references [1] p. bouchner, r. pieknik, s. novotny, et al. fatigue of car drivers – detection and classification based on the experiments on car simulators. 2006, wseas transactions on systems, 5 (12), pp. 2789-2794. [2] p. bouchner. a complex analysis of the driver behavior from simulated driving focused on fatigue detection classification. 2006, wseas transactions on systems, 5 (1), pp. 84-91. [3] p.bouchner, m.hajny, s.novotny, et al. car simulation and virtual environments for investigation of driver behavior. 2005, neural network world, 15 (2), pp. 149-163. 40 vol. 12/2017 analysis of stored data help to propose and generate new tracks [4] p. spurny, j. andrs, p. bouchner, et al. testing a system for predicting microsleep. 2016 lekar a technika, 46 (2), pp. 51-54. [5] p. bouchner, s. novotny. system with driving simulation device for hmi measurements. 2005, wseas transactions on systems, 4 (7), pp. 1058-1063. [6] m. matowicki, o. pribyl, p. bouchner. pragmatic overview of surrounding traffic implementation into driving simulator. elektro 2016 – 11th international conference, proceedings, art. no. 7512111, pp. 423-428. [7] h. klee. microscopic car modeling for intelligent traffic and scenario generation in the ucf driving simulator. [8] p. bouchner. driving simulators for hmi research. phd. thesis, ctu. [9] r. pieknik. the methodology of creating scenario roads of driving simulator for different types of experiments. conference of driver-car interaction and interface 2009. [10] a. orlicky. automatic generation of road infrastructure in 3d for vehicle simulators. diploma thesis, prague 2016. [11] j. togelius, r. de nardi, s. lucas. making racing fun through player modeling and track evolution. [12] d. loicono, l. caramone, p. lanzi. automatic track generation for high-end racing games using evolutionary computation. 41 acta polytechnica ctu proceedings 12:38–41, 2017 1 introduction 2 measured data analysis 3 identifying common elements of behavior 4 proposal track by type of experiment 5 the algorithm for automatic generation of experimental tracks 6 pseudo-code 7 conclusions references 246 acta polytechnica ctu proceedings 2(1): 246–251, 2015 246 doi: 10.14311/app.2015.02.0246 recurrent novae — a review k. mukai1,2 1cresst and x-ray astrophysics laboratory, nasa/goddard space flight center, greenbelt, md 20771, usa 2department of physics, university of maryland, baltimore county, 1000 hilltop circle, baltimore, md 21250, usa corresponding author: koji.mukai@nasa.gov abstract in recent years, recurrent nova eruptions are often observed very intensely in wide range of wavelengths from radio to optical to x-rays. here i present selected highlights from recent multi-wavelength observations. the enigma of t pyx is at the heart of this paper. while our current understanding of cv and symbiotic star evolution can explain why certain subset of recurrent novae have high accretion rate, that of t pyx must be greatly elevated compared to the evolutionary mean. at the same time, we have extensive data to be able to estimate how the nova envelope was ejected in t pyx, and it turns to be a rather complex tale. one suspects that envelope ejection in recurrent and classical novae in general is more complicated than the textbook descriptions. at the end of the review, i will speculate that these two may be connected. keywords: cataclysmic variables symbiotic stars recurrent novae individual: t pyx. 1 introduction nova eruptions are understood to be powered by thermonuclear runaway (tnr) on the surface of accreting white dwarfs. hundreds of objects in the galaxy have been seen to experience one nova eruption: these are called classical novae (cne). recurrent novae (rne) are objects that have been seen to experience multiple nova eruptions. there are currently 10 confirmed rne in the galaxy. between 10−6 and 10−4 m� of hydrogen rich material needs to be accreted to reach the critical temperature and density required for tnr. the critical mass is lower for more massive white dwarfs with higher gravity. therefore, we expect rne to contain near chandrasekhar mass white dwarf accreting at a high rate. this makes rne candidate progenitors of type ia supernova. for this reason, and because the recurrent nature of these objects allows studies that one cannot undertake for cne, rne have become the subject of intensive study. it is impossible to present a comprehensive review of rne in the space allotted; for that, the readers are referred to schaefer (2010) and anupama (2013). in this review, i will present selected highlights from multiwavelength campaigns on recent rn outbursts, highlighting the work of the e-nova collaboration1. i will also include results on several cne: some of these system may be unrecognized or unconfirmed rne, and others provide a useful comparison. i will also present some quiescent observations. i will discuss implications on the white dwarf mass, the ejecta mass, the quiescent accretion rate,and the evolutionary scenarios for rne and cne. 1.1 x-ray bursts: a cautionary tale although rne provide a unique opportunity to compare multiple nova episodes and possibly to compare accreted vs. ejected mass, only a handful of eruptions are observed for each system. this is in stark contrast to the studies of x-ray bursts, which are tnrs on accreting neutron stars. for example, linares et al. (2012) studied 398 x-ray bursts detected from the transient x-ray binary in the globular cluster, terzan 5, as the accretion changed by a factor of ∼5. this allowed these authors to study the relationship between the persistent luminosity, the burst recurrence time and the burst fluence, and thereby test the theory of tnr on neutron stars. unfortunately, analogous tests have not been possible yet in the case of rne. yet, even in the case of x-ray bursts, puzzles remain (galloway et al. 2008). one is the burst oscillations observed during the decay. the drifting period of burst oscillations reflect the spin period of the neutron star atmosphere, which changes as the atmosphere expands and then contracts during the course of a burst. the presence of the oscillations during the decay, however, requires inhomogeneous burning over the neutron star 1https://sites.google.com/site/enovacollab/ 246 http://dx.doi.org/10.14311/app.2015.02.0246 recurrent novae — a review surface, even though one might expect uniform burning at this stage. the other is that pairs of bursts can occur with very short (<10 min) recurrence times, much too short to have accreted sufficient fuel for a new burst, judging by the persistent x-ray luminosity. this requires a reservoir of unburnt fuel on or very near the neutron star surface. thus, our theoretical understanding of x-ray bursts appears incomplete. it may well be that the current theories of nova outbursts are equally incomplete regarding, e.g., the recurrence times of rne. 2 selected recent results 2.1 ejecta geometry montez et al. (in preparation) have detected extended x-ray emission in the chandra observations of rs oph obtained in 2009 and 2011. these structures are wellseparated from the central x-ray source in the e-w direction, and were seen to expand from 2009 to 2011. this x-ray emitting bipolar outflow appears to follow the same angular expansion curve inferred for radio and hubble space telescope (hst) bipolar structures observed earlier. the implied current expansion velocity is very high (of order 4,000 km s−1. one possible origin of the bipolar flow is that rs oph produced a true, well-collimated, jet near the time of nova eruption. another is that an initially spherical ejecta encountered an equatorial torus and slowed down except in the polar directions. since rs oph is an rn in a symbiotic binary, the wind of the giant mass donor is a potential source of such a torus (mohamed et al. 2013). however, similar shaping of the ejecta might also occur in cataclysmic variables (cvs), with a roche-lobe filling mass donor on or near the main sequence. in a series of simulations of the 2010 eruption of u sco by drake & orlando (2010), the accretion disk is destroyed by the blast wave. this interaction causes the ejecta to expand away from the orbital plane. one particular simplifying assumption used by these authors, that of a uniform density disk, is a cause for concern, and independent simulations are needed to confirm their results in general. nevertheless, the possibility that disk-blast wave interactions create bipolar outflow should be kept in mind for all novae, whether the underlying binary is a symbiotic system or a cv. the above-mentioned results on rs oph and u sco are both about the outflow during the most recent outbursts of rne, and may apply to cne as well. in contrast, one type of study unique to rne is the analysis of light echoes produced by ejecta from previous outbursts, as sokoloski et al. (2013) did for t pyx. the arrangement of the echo location on the sky and the progression of echos from east to west suggest a ringlike structure from a previous outburst. the delay times for echoes along the north-south axis suggest a distance of 4.8±0.5 kpc for t pyx. moreover, the time lags between different echoes suggest that the ring is inclined ∼30–40◦ relative to the plane of the sky. this is most likely to reflect the binary inclination, somewhat higher than values previously inferred for t pyx. regardless of the precise inclination angle, the very fact that an equatorial ring was formed by the ejecta is worth noting. 2.2 novae in symbiotic systems four of the known galactic recurrent novae are in symbiotic binaries: rs oph, t crb, v745 sco, and v3890 sgr. they are all s type systems: they have a normal red giant mass donor, an orbital separation of order 1 au, and an orbital period of order 1–2 years. in the other subtype, the d (dusty) type, the mass donor is an agb star; the d type systems have a much wider orbit than the s type systems. before 2010, all known tnr events in symbiotic systems were either a very slow “symbiotic novae,” or very fast rne (mikolajewska 2008). it is important to note that tnr can lead to a quasi-static configuration without explosive mass loss in symbiotic novae. also noteworthy is the fact that the accretion rate is high enough in the 4 s type symbiotic systems to produce rne. if we take 10−7 m� yr −1 as the typical wind mass loss rate of a normal red giant, then this implies either roche-lobe overflow or a very efficient mechanism to capture the wind, such as wind roch-lobe overflow (mohamed & podiadlowski 2010), although m giants in symbiotic binaries may have higher mass-loss rates (seaquist & taylor 1990). in march 2010, a d-type symbiotic system, v407 cyg, became a nova. it was noteworthy for being the first nova to be detected as gev γ-ray source with fermi lat (abdo et al. 2010); it was the subject of an intensive multiwavelength from radio to x-rays (nelson et al. 2012; chomiuk et al. 2012). the x-rays were predominantly from the shock between the nova blast wave and the wind of the mira type mass donor; interestingly, v407 cyg became x-ray bright after the gev signal faded. the thermal emission from the flashionized agb wind was the dominant source of radio signal. while we learned a lot about the nova event, we are left with one important question: how often does v407 cyg experience nova eruptions? is it an unrecognized rn, or are the eruptions much less frequent? 2.3 long period cvs darnley et al. (2012) proposed to classify novae into red giant, sub-giant, and main sequence systems. the orbital periods of the “sub-giant” systems are in the range 10 hrs to 6 days. according to the ritter & kolb catalog version 7.20 (ritter & kolb 2003), there are 46 247 k. mukai systems (excluding one uncertain entry in the catalog) in this orbital period range. of the confirmed rne, v894 cra, u sco and ci aql belong to this group, and a fourth, v2487 oph, probably has an orbital period in this range. the evolution of such long-period cvs has not been studied extensively to date, compared to those with periods under 10 hrs, for which the basic framework and much more have been established (knigge et al. 2011). cvs with similarly long orbital periods include several novae not known to be recurrent (gk per, v1017 sgr), supersoft sources (e.g., mr vel, cal 83), and v sge and other systems that may be related to supersoft sources. one possibility is that these systems are currently undergoing thermal timescale mass transfer (ttmt; schenker et al. 2002) or did so in the past. the evolution of cvs with p>10 hrs should be studied further, and a search for additional nova outbursts of systems like gk per may be worthwhile. 2.4 quiescent x-ray observations figure 1: possible white dwarf mass of v2487 oph for the magnetic (dashed line) and the non-magnetic (solid line) cases of the known recurrent novae, t crb and v2487 oph are bright hard x-ray sources detected with integral and with swift bat (see, e.g., baumgartner et al. 2013). when such hard x-ray emission is detected, it can constrain both the white dwarf mass and the accretion rate in quiescence. in both magnetic and nonmagnetic cases, hard x-rays are generated when the supersonic accretion flow encounters the white dwarf surface and is shock-heated. the hot plasma must radiatively cool before settling onto the white dwarf surface. in the case of magnetic cvs, the accretion flow is radial and its velocity is approximately at the freefall value (aizu 1973). the observed x-ray spectrum is multi-temperature in nature, with shock temperatures typically in the 10–50 kev range. the spectral curvature in the hard x-ray range is a reliable measure of the shock temperature, and can be used to infer the white dwarf mass (see, e.g., yuasa et al. 2010). the situation may well be more complex for the boundary layer of non-magnetic cvs, but the maximum temperature is unlikely to exceed that expected for strong shocks from keplerian velocity, and quiescent dwarf novae appear to follow the “keplerian strong shock” relationship (byckling et al. 2010). some authors have speculated that v2487 oph may contain a magnetic white dwarf, based solely on its bright, hard x-ray emission. however, i measure a shock temperature of ∼50 kev using the bat survey 8-channel spectrum, which corresponds to either a ∼0.9 m� magnetic white dwarf or a >1.24 m� non-magnetic white dwarf. given the rn nature of this object, and given that repeated xmm-newton observations have not revealed a spin modulation, the latter interpretation seems more likely. 2.5 high mass transfer rate: evolutionary or temporary? patterson et al. (2013) made the following simple prediction, based on average secular mass accretion rate: cvs above the period gap should experience nova eruptions once every 10,000 years or so, while those below the gap should recur once every 1 million year. for these normal cvs to be a rn, the accretion rate needs to be elevated by many orders of magnitudes above the secular mean. this is in contrast to the long-period systems with a sub-giant mass donor, where ttmt may drive a very high accretion rate. similarly, in symbiotic systems, the mass loss rate from the donor is high enough, although the fraction that can be captured by the white dwarf is highly uncertain. let us now consider the census of known classical and recurrent novae of various types with the above expectations in mind. among symbiotic stars, most novae are extremely slow and often referred to as symbiotic novae, the slowness suggestive of low-mass white dwarfs. the four wellknown symbiotic recurrent novae are all in s-type systems: the short recurrence time and the high velocity of the ejecta both suggest these to have massive white dwarfs. the bifurcation of novae in symbiotic stars into two such extreme groups is very different from the situation in cvs, and should be investigated further. although only a single outburst is known, the outburst properties of v407 cyg makes it a candidate rn in this context. since it is in a d type symbiotic system, with a much greater binary separation, an estimate of its quiescent accretion rate, when one is obtained, may tell us about how white dwarfs accrete in symbiotic stars, not to mention providing a clue as to its likely recurrence time. u sco, ci aql, and v894 cra are long-period systems; v2487 oph may belong in this group, although its orbital period has not been determined yet. the ttmt 248 recurrent novae — a review scenario points to the general framework for why these systems can be rne. more research is needed to understand why some long-period systems are sss, others rne, and yet others cne (as far as we know). this leaves im nor (in the period gap) and t pyx (below the gap) in the orbital period range with very low secular accretion rate and with only a small number of known classical novae (v1794 cyg, v per, ...). it is interesting to note that, while many classical novae are known above the gap (orbital period in the 3–10 hr range), no rne are known in this regime. this could purely a matter of small number statistics. at the same time, it could be that the mechanism elevating the accretion rate of t pyx and im nor far above the secular mean is much less effective for systems above the gap. 2.6 multi-wavelength observations of t pyx multi-frequency radio monitoring of novae is a powerful technique that allows us to estimate the total ejected mass in a relatively simple manner, although complications often arise (see, e.g., roy et al. 2012 and references therein.) the nova ejecta is initially optically thick in the radio, thus the brightening traces the angular expansion of the ejecta. as the ejecta becomes optically thin, first from the highest frequencies then progressively to lower frequencies, this allows the amount of mass to be estimated, as long as we have a handle on the temperature and clumping of the ejecta. at the same time, we know that early x-ray emission from novae is likely due to shocks in the ejecta (e.g., o’brien et al. 1994). with this mind, the e-nova collaboration has begun multi-wavelength observations of recent novae using the much improved karl g. jansky vla in the radio, and swift and other observatories in the x-rays. t pyx is one of the major targets of the e-nova collaboration. the radio and x-ray results are presented by nelson et al. (2014) and chomiuk et al. (2014), respectively. t pyx was largely undetected in the radio for the first ∼60 days since the discovery of the 2011 outburst, then started to rise around day ∼100. it was also x-ray faint during the first several months, and then started to rise slightly after the onset of the radio rise. the x-ray photons detected with swift xrt are a mixture of optically thin emission from the shocked shell and the supersoft emission from the still nuclear burning surface. it is difficult to disentangle the two using short snapshot observations, hence it is difficult to determine, e.g., the turn-on time of the supersoft emission. in the optical, t pyx remained near peak optical magnitude for 2 or 3 months, depending on where one defines the peak to have ended and decline to have begun. this implies a large photospheric radius, perhaps of order 5×107 km (∼ a third of an au; assuming a blackbody with a temperature of 10,000k, a distance of 4.8 kpc, and av ∼ 1.0, this radius corresponds to a visual magnitude of v∼7.9). this is well outside the central binary, yet it only takes of order 1 day for matter traveling at 600 km s−1 to reach this distance. for a shell at a distance of 1×108 km to have an optical depth of 1, which implies a column density of order 3 g cm−2 (assuming electron scattering dominates the opacity), the total mass of the shell must be greater than ∼ 5.0 × 10−7 m�. so the duration of optical peak implies either a continuous ejection of ∼ 5×10−7 m� per day for several months, or that t pyx went into a quasi-static, red giant-like configuration during the peak, and ejected the extended atmosphere with a significant delay. in the latter case, the total mass of the extended atmosphere must be much greater than 5×10−7 m�, because it is a filled sphere and not a thin shell. schaefer et al. (2013) presented the detailed visual light curve of t pyx during the initial rise. until it reached v∼7.7, it can be modeled well assuming a uniform expansion of the ejecta, then the observed brightness drops below this model. if we equate this instance with the time when the optical depth of the ejecta dropped below 1.0, we infer that an initial shell of ∼ 6 × 10−7 m� was ejected. such a small ejection is easy to hide in the radio data, although there is one detection on day 17 that could be interpreted as due to this. on the other hand, continuous mass ejection of ∼ 5 × 10−7 m� per day is difficult to reconcile with the deep radio non-detections followed by rapid brightening around day 60. rather, the radio data are consistent with a prolonged period of quasi-stationary configuration, and a delayed ejection of a more massive (of order 10−5 m�) shell. in addition, if the latter system was ejected with a larger velocity, then we expect shocked x-ray emission when it catches up with the initial ejecta. the existing x-ray data are broadly consistent with such a picture. 2.7 the cause of elevated mass transfer rate for t pyx, we have a clear-cut case that the mass transfer rate is elevated by several order of magnitudes above the evolutionary mean (gilmozzi & selvelli 2007). the same presumably applies to im nor as well. irradiation of the donor is often invoked as the explanation. however, theorists have long concluded that this requires hard photons, and therefore theoretical studies largely focus on x-ray binaries (see, e.g., hameury et al. 1986; king 1989). to quote from ritter (2000), “energy emitted in certain spectral ranges, 249 k. mukai as e.g., euv radiation and soft x-rays, is unlikely to reach the photosphere of the donor.” to reach the photosphere of k or m type dwarfs, irradiating flux needs to be able to penetrate above 1024 cm−2 of column, thus requiring strong flux above 10 kev. therefore, supersoft sources or photospheric emission of otherwise very hot white dwarf only irradiates the chromosphere and above, and not the photosphere. the irradiation mechanism studied in above-mentioned papers cannot work when the irradiating flux is in the form of soft x-ray and euv photons. while v1500 cyg is sometimes taken as an example of a system that is experiencing enhanced accretion rate due to irradiation, this is not necessarily the case. this is because one well-established effect of irradiation is to increase the luminosity of the existing structures, be it the secondary or the accretion disk. in fact, according to somers & naylor (1999), the elevated brightness of v1500 cyg, which is currently an asynchronous polar due to its 1975 nova eruption, is due to an orbitally modulated component, not due to the spin modulated component. since accretion luminosity should be modulated on the spin period, we know that the extra light is due to the irradiated face of the secondary. in this picture, the reflection of the gradually decreasing postnova white dwarf flux explains the secular changes in the brightness of v1500 cyg, without invoking variable accretion rate. thomas et al. (2008) obtained phase-resolved kband photometry of old novae of various ages since outburst. they were also able to interpret their results without invoking changing accretion rate. rather, in their interpretation, variable irradiation, and hence variable reflection, changes the brightness of the existing structures, the accretion disk and the secondary. at a minimum, this points out that an enhanced brightness is insufficient to prove an enhanced accretion rate. these studies also suggest that irradiation by a postnova does not lead to enhanced mass transfer, although they do not yet constitute a solid proof. if not irradiation, what other mechanisms can potentially enhance the mass transfer? here i speculate that the post-nova common envelope phase might be ultimately responsible, as follows. the ring geometry of the ejecta of t pyx (§2.1) suggests that the secondary plays a role in shaping the geometry of the ejecta. in symbiotic systems, slow novae can stay in the “plateau” phase for decades, whereas no such example is known among cvs, again implying that the secondary plays a role in ejecting the nova envelope. the potential role played by the common envelope system was first pointed out by macdonald (1980) and later studied quantitatively by, e.g., livio et al. (1990). the general consensus is that the common envelope phase can contribute to the ejection of the nova envelope, but only if the ejecta is moving more slowly than the orbital motion of the mass donor. the multi-wavelength data on t pyx indeed suggests that the envelope may have been in a quasi-stationary configuration for 2–3 months, as it is for decades in slow symbiotic nova. the possibility that many novae in cvs may not be able to eject the bulk of the envelope, if it were not for the common-envelope phase, should therefore be investigated. if common envelope phase is an important factor in the envelope ejection process, then this implies that the binary must have lost angular momentum. while the orbital period of t pyx was seen to increase after the 2010 eruption (patterson et al., this volume), the period increase only constrains the combination of ejected mass and angular momentum loss. if more than the minimum allowed mass was ejected, so was angular momentum. so the proposed scenario is that of a common-envelope interaction during the nova eruption, resulting in an impulsive angular momentum loss, which drives a higher accretion rate for the ensuing decades. i suggest that such a scenario deserves serious, quantitative analysis. 3 summary and conclusions recent rn eruptions have be subjected to intense multi-wavelength observing campaigns using advanced facilities, including the karl g. jansky vla, many ground-based optical/ir telescopes, hst, swift and other x-ray observatories. although not a new discovery as such, recent images of nova ejecta demonstrates once again that they are not spherically symmetric. in the case of t pyx, the multi-wavelength data strongly suggest that there was a initial, small ejection and a much larger, delayed ejection. these facts together suggest that the binary companion, via common-envelope interaction, may be involved in the ejection process. it is expected that rne harbor massive white dwarfs accreting at a very high rate. while the accretion rate may be high in some subset of rne for evolutionary reasons (symbiotic rne, and long-period systems), this is definitely not the case for t pyx and im nor. unfortunately, there is a serious difficulty with the commonly invoked mechanism of irradiation, when the irradiation is by soft x-ray and euv photons. i presented a possible scenario involving impulsive angular momentum loss during the common envelope phase. acknowledgement i thank my colleagues in the e-nova collaboration, particularly drs. nelson, sokoloski, and chomiuk, for stimulating discussion. 250 recurrent novae — a review references [1] abdo, a.a. et al.: 2010, sci. 329, 817. doi:10.1126/science.1192537 [2] aizu, k.: 1973, prog theor. phys. 49, 1184. doi:10.1143/ptp.49.1184 [3] anupama, g.g.: 2013, iausymp 281, 154. 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[35] yuasa, t. et al.: 2010, a&a 520, a25 discussion christian knigge: regarding the possibility that irradiation might enhance the mass accretion rate: i think that an enhancement is possible, at least via irradiation-driven winds from the donor. in fact, this model was first proposed for sss, where the irradiation is all in the soft x-ray/euv regime. koji mukai: it is true that the viability of irradiation-driven wind to enhance the accretion rate is a separate issue. i feel that sometimes the direct effect of irradiation is often too casually invoked as the explanation, and my intent was to advise caution. 251 http://dx.doi.org/10.1126/science.1192537 http://dx.doi.org/10.1143/ptp.49.1184 http://dx.doi.org/10.1088/0004-637x/761/2/173 http://dx.doi.org/10.1088/0004-637x/788/2/130 http://dx.doi.org/10.1088/0004-637x/746/1/61 http://dx.doi.org/10.1088/2041-8205/720/2/l195 http://dx.doi.org/10.1086/592044 http://dx.doi.org/10.1093/mnras/241.3.365 http://dx.doi.org/10.1088/0067-0049/194/2/28 http://dx.doi.org/10.1088/0004-637x/748/2/82 http://dx.doi.org/10.1086/168836 http://dx.doi.org/10.1088/0067-0049/187/2/275 http://dx.doi.org/10.1088/0004-637x/773/1/55 http://dx.doi.org/10.1046/j.1365-8711.2002.05999.x http://dx.doi.org/10.1088/2041-8205/770/2/l33 introduction x-ray bursts: a cautionary tale selected recent results ejecta geometry novae in symbiotic systems long period cvs quiescent x-ray observations high mass transfer rate: evolutionary or temporary? multi-wavelength observations of t pyx the cause of elevated mass transfer rate summary and conclusions 66 acta polytechnica ctu proceedings 2(1): 66–70, 2015 66 doi: 10.14311/app.2015.02.0066 multifrequency behaviour of polars k. reinsch1 1georg-august-universität göttingen, institut für astrophysik, friedrich-hund-platz 1, 37077 göttingen, germany corresponding author: reinsch@astro.physik.uni-goettingen.de abstract cataclysmic variables emit over a wide range of the electromagnetic spectrum. in this paper i will review observations of polars in relevant passbands obtained during the last decade and will discuss their diagnostical potential to access the physics of the main components within the binary systems. this will include a discussion of intrinsic source variability and the quest for simultaneous multi-frequency observations. keywords: cataclysmic variables polars optical spectroscopy photometry ir uv x-rays. 1 introduction in the standard model of magnetic cataclysmic variables a polar is described as a semi-detached binary in which matter from a roche-lobe filling late-type mainsequence star is coupled outside the circularization radius to the field lines of a highly-magnetic white dwarf and channeled onto small accretion region(s) in the vicinity of the magnetic pole(s). the system components can be observationally disentangled – at least to some degree – by their characteristic emission in different wavelength regimes (see fig. 1). the contribution of the binary components to the spectral energy distribution of the system will be discussed individually and in the context of multi-frequency observations. the basic understanding of polars has not much changed since the comprehensive review by warner (1995). i will, therefore, concentrate on recent observational findings which led to a more detailed understanding of the physical processes involved. figure 1: schematic view of a polar and main emission regions (iris traulsen, adapted from marc garlick). 66 http://dx.doi.org/10.14311/app.2015.02.0066 multifrequency behaviour of polars 2 binary components and spectral energy distribution 2.1 secondaries in polars secondaries in polars emit most of their light in the red and near ir. but even in these bands their luminosity is generally outshone by other components within the binary. the secondaries are, therefore, best studied when the system happens to be in a low state of accretion or during eclipses of the accretion region and the white dwarf. optical and near-ir studies have revealed that their spectra are consistent with normal late-type dwarfs (or brown dwarfs). contrary to similar studies of non-magnetic systems, no significant abundance anomalies have been found from near-ir spectroscopy of the secondaries in polars. this provides evidence that magnetic and non-magnetic systems might take different evolutionary paths (harrison et al., 2005). spectroscopy during the eclipse of the primary combined with phase-resolved near-ir photometry of polars can be used to determine the spectral type and the brightness of the secondary, as has been shown e.g. for the long-period polar v1309 ori (reinsch at al., 2006). using calibrations of the surface brightness of m-dwarfs (e.g. beuermann 2000), from the brightness of the secondary an accurate distance of the system can be derived. 2.2 circumbinary dust disk mid-infrared observations of the short-period polar ef eri with the infrared spectrograph (irs) on the spitzer space telescope have revealed the presence of emission in the 3–14 µm wavelength range which exceeds the combined spectral energy distribution of the white dwarf, the l5-type brown dwarf secondary, and cyclotron radiation from the b ≈ 13 mg white dwarf magnetic field. hoard et al. (2007) concluded that the spectrum longward of ∼ 5 µm is dominated by emission from a circumbinary dust disk. in a later paper, however, harrison et al. (2013) argue based on combined ground-based, spitzer, wise, and herschel observation of the two polars am her and ef eri that their midinfrared excess can be fully explained by optically thin emission from the lowest cyclotron harmonics (n ≤ 3). 2.3 stellar activity of the secondary from long-term observations of polars it is well known that most systems alternate between low and high states and sometimes also intermediate states of accretion. the best studied system is am her with a record of 33 years of optical observations by the american association of variable star observers (aavso, http://www.aavso.org/). due to the absence of an accretion disk in polars such changes must reflect the immediate response of the system on variations of the mass-transfer rate. stellar activity of the donor star has been proposed as a possible mechanism to induce such variations (e.g. hessman et al., 2000). in a recent analysis of shortand long-term brightness variations of five polars, periodic variations were indicated which could result from magnetic cycles in the secondary stars (kalomeni, 2012). kafka et al. (2010) present another piece of evidence for stellar activity from phase-resolved high-resolution spectroscopy of polars in low states. they interprete two bright satellites of the central low-velocity component in the hα doppler map of bl hyi as prominencelike magnetic loops kept in place by magnetic field interactions between the white dwarf and the donor star. howell et al. (2006) have monitored the long-term variation of the hα equivalent width of ef eri during its extended low state and attribute the observed behaviour to chromospheric activity of the secondary. in addition they identify magnetic field interaction between the two stars as a probable mechanism that would concentrate stellar activity on the white dwarf facing hemisphere of the secondary. 2.4 coupling region partially ionized matter released from the secondary over the l1 point is assumed to follow initially a freefalling trajectory. at the magnetospheric radius of the white dwarf the ram pressure of the infalling gas is balanced by the magnetic pressure and the accretion stream couples to the magnetic field of the white dwarf. observationally little information is available on the processes occurring in this coupling region. xmm-newton observations of v1309 ori show a prolonged ingress of the soft x-ray light curve with xray emission being visible beyond the geometrical occultation of the white dwarf. schwarz et al. (2005) have suggested plasma heated by shocks in the coupling region as a possible origin of the x-ray emission seen during eclipse. 2.5 accretion curtain beyond the coupling region accretion becomes magnetically controlled and the initially narrow ballistic stream is diverted off the orbital plane and forms a broad curtain-like structure through which matter is channeled into the vicinity of the magnetic poles of the white dwarf. the accretion stream and curtain are characterized by an optical and uv emission-line spectrum comprising strong lines of neutral hydrogen and of ionized helium. using relatively simple geometrical assumptions indirect imaging techniques, like doppler 67 k. reinsch tomography, roche tomography, accretion stream mapping, and eclipse mapping, have been well established to confine the geometrical parameters of different parts of the accretion stream (e.g. schwope et al., 2004). 2.6 white dwarf white dwarfs in polars have typical temperatures in the range 10,000–15,000 k. their photospheric emission has been well studied in the uv and far-uv spectra of polars in low states of accretion. based on a far-uv spectroscopic survey with hst/ stis, araujo-betancor et al. (2005) found that at any given orbital period white dwarfs in magnetic cvs are colder than those in non magnetic cvs with the differences in effective temperatures being largest for systems above the period gap. as a possible explanation they suggest that magnetic braking in polars might be reduced. the magnetic field structure of several white dwarfs in polars has been studied in detail using phase-resolved circular spectropolarimetry and zeeman tomography (e.g. beuermann et al., 2007 and references therein). in most cases the field structure departs significantly from that of a dipole field and higher order multipoles are required to adequately describe the observed zeeman spectra. accreting material is flowing preferentially along field lines reaching far out towards the ballistic part of the stream. depending on field geometry and accretion rate, this leads to one or more accretion regions on the white dwarf. 2.7 accretion region in the standard scenario actual accretion onto magnetic white dwarfs occurs in a comparably small region of some 0.1% of the white dwarf surface. infalling matter forms an accretion column with a shock front at some height above the white dwarf where the supersonic gas of the accretion stream is decelerated to a subsonic settling flow. the shock height and the depression of the post-shock flow below the photosphere depends on the magnetic field strength, the mass of the white dwarf, and the specific accretion rate. the structure of the accretion column has been studied theoretically. the emergent spectra have been calculated by radiationhydrodynamical models as a function of mass-flow rate and field strength (e.g. fischer & beuermann, 2001; beuermann, 2004). at low mass-flow rates, the shock stands above the atmosphere and bremsstrahlung radiation is the main cooling process. for larger accretion rates the emission is hidden behind the column density and reprocessed as soft x-rays. irradiation is the principle cause for a large heated accretion cap extending far beyond the impact region of the accreted material. calculations with a standard lte stellar atmosphere code including an angledependent external radiation source have demonstrated that much of the reprocessed light appears in far uv and not at soft x-rays (könig et al., 2006). observationally the temperature distribution in the accretion spot could be demonstrated by the analysis of the chandra letg spectrum of am her. this data provide enough photons to show that the soft x-ray component cannot be fitted by a single temperature model (beuermann et al., 2012). at high mass-transfer rates sufficiently dense blobs can penetrate to large optical depths in the white-dwarf atmosphere submerging the shock below the photosphere. the most extreme example of this “blobby accretion” case has been observed in v1309 ori. its soft x-ray light curve obtained with xmm-newton shows x-ray emission occuring in flares of typically 10 s time scale. this corresponds to a length scale of 4 × 109 cm for the infalling filaments (schwarz et al., 2005). short-term spectral variability at x-ray wavelengths could be studied in the rosat pspc high state observation of am her which contains 1.3 million counts. significant variability on time scales down to 200 ms could be detected. correlated spectral and count-rate variations are seen in flares on time scales down to 1 s, demonstrating the heating and cooling associated with individual accretion events (beuermann et al., 2008). some polars are even bright enough that x-ray emission line spectroscopy is feasible with current instruments. the high-diagnostics potential of such observations has been demonstrated e.g. by girich et al. (2007), mauche et al. (2003), mauche et al. (2009). 3 variability of spectral energy distribution several polars have been studied at different epochs yielding that their optical, uv, and x-ray fluxes can vary by more than an order of magnitude. as an example, fig. 2 shows the integrated energy flux νfν from the infrared to the hard x-ray regime of the polar rx j1007.5-2017 collected with various instruments between 1992 and 2001. the left panel shows spectrophotometry of 1992, 1997, and 2000 (black curves), supplemented by the 2mass j, h, and k-band fluxes of 1999 (red filled squares) and the visual and ultraviolet fluxes measured with the optical monitor on board of xmm-newton simultaneous to the 2001 x-ray observation (red filled circles). the green curve represents the adjusted flux distribution of the dm3 star lhs58 which is expected to match the spectral type of the donor star and to be the main contribution to the 1997 low-state spectrum. 68 multifrequency behaviour of polars figure 2: spectral energy distribution of the polar rx j1007.5-2017 at different epochs (from: thomas et al. 2012) the right panel shows the incident spectra for the 2001 xmm pn observation (red filled circles) and the 1992 rosat pspc observation (open circles). the dashed curve illustrates the “source spectrum” corrected for interstellar absorption (thomas et al., 2012). 4 conclusions x-ray and optical surveys, like the rosat all-sky survey and the sloan digital sky survey, let to the discovery of a large number of new polars during the recent two decades increasing the sample size to 135 polars with known orbital periods (ritter & kolb 2003; 2013). in addition, sensitivity and instrumental capabilities at all wavelengths have been significantly enhanced. taking both together, polars provide an important lab to quantitatively understand individual systems and physical processes of matter under extreme conditions. in this sense, we are indeed currently in a “golden age” of research on polars. there are still several fundamental questions remaining open. among these are: do we fully understand the evolution of polars? what drives accretionrate variations? do polars with fields as high as in single white dwarfs exist? what is the origin of the magnetic field in polars? what exactly happens in the coupling region? multi-frequency observations are a powerful tool to improve our understanding of polars in future. due to large variations of the accretion rate on different time scales simultaneous coverage at different wavelengths is mandatory. acknowledgement i thank the many people who contributed to increase our knowledge on polars over the last decade, and kindly acknowledge fruitful and long-lasting collaborations with klaus beuermann, vadim burwitz, essam el-kholy, fabian euchner, boris gänsicke, yonggi kim, axel schwope, robert schwarz, hans-christoph thomas, iris traulsen, and fred walter which led to several papers i have referred to in this contribution. references [1] araujo-betancor, s., gänsicke, b.t., long, k.s., beuermann, k., de martino, d., sion, e.m., szkody, p., 2005, apj 622, 589 doi:10.1086/427914 [2] beuermann, k., 2000, new astr. rev. 44, 93 doi:10.1016/s1387-6473(00)00020-8 [3] beuermann, k., 2004, aspc 315, 187 [4] beuermann, k., el kholy, e., reinsch, k., 2008, a&a 481, 771 [5] beuermann, k., euchner, f., reinsch, k., gänsicke, b.t., 2007, a&a 463, 647 [6] beuermann, k., burwitz, v., reinsch, k., 2012, a&a 543a, 41 [7] fischer, a., beuermann, k., 2001, a&a 373, 211; 69 http://dx.doi.org/10.1086/427914 http://dx.doi.org/10.1016/s1387-6473(00)00020-8 k. reinsch [8] girich, v., rana, v.r., singh, k.p., 2007, apj 658, 525 [9] harrison, t.e., howell, s.b., szkody, p., cordova, f.a., 2005, apj 632, l123 doi:10.1086/498067 [10] harrison, t.e., hamilton, r.t., tappert, c., hoffman, d.i., campbell, r.k., 2013, aj 145, 19. [11] hessman, f., gänsicke, b., mattei, j., 2000, a&a 361, 952 [12] hoard, d.w., howell, s.b., brinkworth, c.s., ciardi, d.r., wachter, s., 2007, apj 671, 734. doi:10.1086/522694 [13] howell, s.b., walter, f.m., harrison, t.e., huber, m.e., becker, r.h., white, r.l., 2006, apj 652, 709 doi:10.1086/507603 [14] kafka, s., tappert, c., ribeiro, t., honeycutt, r.k., hoard, d.w., saar, s., 2010, apj 721, 1714 doi:10.1088/0004-637x/721/2/1714 [15] könig, m., beuermann, k., gänsicke, b., 2006, a&a 449, 1129 [16] mauche, c.w., liedahl, d.a., fournier, k.b., 2003, apj 588, l101 doi:10.1086/375684 [17] mauche, c.w., 2009, apj 706, 130 doi:10.1088/0004-637x/706/1/130 [18] reinsch, k., kim, y., beuermann, k., 2006, a&a 457, 1043 [19] ritter, h., kolb, u., 2003, a&a 404, 301 (update rkcat7.20, 2013) [20] schwarz, r., reinsch, k., beuermann, k., burwitz, v., 2005, a&a 442,271 [21] schwope, a., staude, a., vogel, j., schwarz, r., 2004, an, 325, 197 [22] thomas, h.-c., beuermann, k., reinsch, k., schwope, a.d., burwitz, v., 2012, a&a 546, a104 [23] warner, b., 1995, cataclysmic variable stars, cambridge university press discussion david buckley: would you care to comment on the fact that a certain fraction of polars have no directly (observed) determined magnetic fields (e.g. from either detection of cyclotron features or spectropolarimetry) and have been suggested to be low-field systems. klaus reinsch: the magnetic field of polars must be strong enough that spin and orbital motion of the white dwarf will be synchronized. direct measurements of magnetic fields in polars require the systems to be in a suitable state that either cylcotron or zeeman features can be detected. therefore, there are still many polars lacking field determinations. augustin skopal: how does the sed look like with the unabsorbed supersoft x-rays and dereddened far-uv fluxes? does your unabsorbed x-ray model fit the far-uv fluxes? klaus reinsch: the “source spectrum” of the x-ray components corrected for interstellar absorption is included as the dashed curve in fig. 2. it is two orders of magnitude below the observed flux in the uv. raymundo baptista: you showed doppler tomograms of hα emission from am her in low state by kafka et al. (2010), with side emission apparently from l4 and l5 regions. an alternative explanation for that is zeeman-doppler spliting produced by the magnetic field of the donor star, of the order of a few kg. klaus reinsch: yes, i agree. linda schmidtobreick’s comment: you showed the prominences detected by stella kafka in am her. we detected the same structure in the low state of bb dor which is a non-magnetic vy scl system. so we think it unlikely that the magnetic field of the wd has an influence on this structure and it should rather be explained by a combination of the magnetic field of the secondary plus the roche potential. klaus reinsch: this is indeed interesting to note. in the final remarks of their paper kafka et al. (2010) state that prominence-like structures are likely present in disk cvs above the period gap as well. but their stability and evolution with time is expected to be different. 70 http://dx.doi.org/10.1086/498067 http://dx.doi.org/10.1086/522694 http://dx.doi.org/10.1086/507603 http://dx.doi.org/10.1088/0004-637x/721/2/1714 http://dx.doi.org/10.1086/375684 http://dx.doi.org/10.1088/0004-637x/706/1/130 introduction binary components and spectral energy distribution secondaries in polars circumbinary dust disk stellar activity of the secondary coupling region accretion curtain white dwarf accretion region variability of spectral energy distribution conclusions 320 acta polytechnica ctu proceedings 1(1): 320–321, 2014 320 doi: 10.14311/app.2014.01.0320 concluding remarks gennady s. bisnovatyi-kogan1 1space research institute, profsoyuznaya str. 84/32, moscow 117997, russia, and national research nuclear university ”mephi”, kashirskoye shosse, 31, moscow 115409, russia corresponding author: gkogan@iki.rssi.ru abstract short comments to the conference talks. keywords: cosmology x-ray sources sne star formation agn jets grb. 1 cosmology sz effect (s. colafrancesco, m. arnaud): does it help investigation of cmb fluctuation? sz influence on the cmb spectrum in clusters interferes with primordial cmb perturbations. so we have: *contamination with cluster scale cmb perturbations *clumpiness of a gas inside cluster does not permit to extract cosmological parameters from the comparison of radio and x ray observations. b. harms: gravitational waves as a dark energy strong gravity branes support gw density constant (seems like vacuum eos), but gw produce gravity, not antigravity, like de. n. panagia: sn1a, wmap, planck are the cosmological parameters coincide inside error bars (hubble constant and de density) ? 1 2 magnetic field in astrophysics *core-collapse sn explosions *cr acceleration in snr *accretion disk structure and coronae formation *jet formation and collimation *magnetic accretion and cyclotron lines *magnetars 2.1 x-ray sources heating of accretion disk corona by magnetic reconnection (s. orlando). about 35 years ago this problem was considered in papers of g. s. bisnovatyi-kogan and s. i. blinnikov ”a hot corona around a black-hole accretion disk as a model for cygnus x-1.” sov.astron.lett. 2, 191-193 (sep.oct. 1976) and a.a. galeev, r. rosner, g.s. vaiana ”structured coronae of accretion disks”. apj, 229, 318-326 (april 1979). in the first paper a convective instability of the radiation dominated region in accretion disks was established, and its connection with formation of a hot corona was considered. ”heat transfer in the region of maximum energy release of an accretion disk will take place mainly by convection, serving to enhance the turbulence and to generate a powerful acoustic flux. the hard x-rays emitted by cyg x-1 (e ≤ 200 kev ) might result from comptonization of soft photons in a corona formed around the disk through this heating.” in the second paper the magnetic mechanism of the corona heating over the convective region of the accretion disk was studied. ”a model for the fluctuating hard component of intense cosmic x-ray sources (such as cyg x-i) is developed, based upon the amplification of magnetic fields by convective motions and differential rotation within a hot (t > 106 k) accretion disk. field reconnection within the inner portion of the disk is shown to be ineffective in limiting field amplification; magnetic fields may therefore attain strengths comparable to the equipartition value, leading to their emergence via buoyancy in the form of looplike structures and resulting in a very hot (t > 108k) magnetically confined, structured corona analogous to the observed structure of the solar corona. the energy balance of these loop structures is examined, and it is shown that the disk soft x-ray luminosity determines the predominant energy loss mechanism in loops: at low disk luminosities, thermal bremsstrahlung from these loops domi1resent discovery (arxiv:1312.3313) of an error in the planck detector at 217 ghz make the differences even less convincing. 320 http://dx.doi.org/10.14311/app.2014.01.0320 concluding remarks nates and contributes a steady, shot-noise-like hard xray component. at high disk luminosities the emerging loops are compton-cooled; the soft x-ray flux from the disk is comptonized by the emerged loops, forming a transient, flarelike hard x-ray component.” schematic drawing of the inner accretion disc corona, heated by magnetic reconnection in the emergent magnetic loops, (from the second paper) is given in fig.1. v. simon: analogy between lmxb ks1731-260 (ms pulsar), and her x-1 (p ≈ 1.24 sec) does not seems to be realistic. in particular, these objects have very different magnetic fields, with 2-4 orders of magnitude difference in strength 2.2 jets, grbs and sgrs a long time before the magnetar concept for sgrs came out, a related object called magnetoid was considered by l.m. ozernoi in the paper ”a theory for the formation and structure of quasistellar radio sources quasar model as supermassive star.” astron. zh. 43, 300-312 (1966). ”a ”magnetoid” is a quasistationary configuration whose equilibrium is governed by a magnetic field, provides an idealized representation of many effects occurring in galactic nuclei and especially in the central parts of quasars. the gravitational energy released through secular contraction of the nucleus is the ultimate source of the intense quasar radiation. the magnetic field is an important intermediary. ... local jets and streams of matter connected with active regions may occur when quasar nuclei reach global quasiequilibrium. although the magnetoid approach furnishes a unified explanation for the basic property of quasistellar radio sources an intense and variable flux emitted over a fairly long period many aspects necessarily remain speculative.” now there is a general opinion that quasars and agn are connected with an accretion into supemassive black holes, may be, besides a minority, represented by w. kundt, who gave a talk ”astrophysics without black holes and extragalactic grbs”. there are some observational evidences which do not support a magnetar model for sgrs, so later it could join its magnetoid predecessor. j. beall: numerical simulations, jets of different scale. heating of the surrounding matter by jet propagation. m. della valle: grb and sne connection questions remain. 2.3 sn 1987a b ashenbach: x ray development of sn 1987a 3 fundamental problems of physics chemistry in astrophysics. susana iglesias groth the talk ”the impact of fullerenos, pahs and amino acides in high energy astrophysics”. thermodynamics and gravity. marco merafina & daniele vitantoni the talk ”data analysis of globular cluster harris catalog in view of the king’s models and evolution” gravitational theory. mariafelicia de laurentis the talk ”testing f(r) theories using the first time derivative of the orbital period of the binary pulsars” leopoldo milano the talk ”eccentric eclipsing detached binaries: a toll for testing gravitational theories” 4 conclusion many topics, covering almost all sides of the modern astrophysics were presented. number of participants and number of reports is increasing. duration of talks is decreasing proportionally from 35 min maximal duration talk (1999), to 25 min (2013) in the conference of japan astronomical society (1978) the duration of each talk was 5 minutes, and some participants managed to show more than 50 transparances during this time (competition). let us hope, that franco giovannelli will find a better solution! thank you franco, and collaborators! 321 cosmology magnetic field in astrophysics x-ray sources jets, grbs and sgrs sn 1987a fundamental problems of physics conclusion 42 acta polytechnica ctu proceedings 1(1): 42–48, 2014 42 doi: 10.14311/app.2014.01.0042 lhc, astrophysics and cosmology giulio auriemma1,2 1università degli studi della basilicata, potenza,italy 2infn sezione di roma, rome, italy corresponding author: giulio.auriemma@cern.ch abstract in this paper we discuss the impact on cosmology of recent results obtained by the lhc (large hadron collider) experiments in the 2011-2012 runs, respectively at √ s = 7 and 8 tev. the capital achievement of lhc in this period has been the discovery of a spin-0 particle with mass 126 gev/c2, very similar to the higgs boson of the standard model of particle physics. less exciting, but not less important, negative results of searches for supersymmetric particles or other exotica in direct production or rare decays are discussed in connection with particles and v.h.e. astronomy searches for dark matter. keywords: hep physics cosmology dark matter dark energy. 1 introduction on july 4th 2012 the atlas[1] and cms[2] collaboration has announced the discovery of a new massive boson, which subsequentely was shown to look like the standard model higgs bosonn. this was possible only after the 2012 runs of lhc at √ s = 8 tev. the two central experiment atlas & cms have collected during 2011 and 2012 an integrated luminosity of ≈ 40 fb−1. several reason support the general belief that this boson could be the long expected higgs boson [3], in particular • it is definitively a boson with spin 6= 1 because decays in γ-γ channel (h → γγ). quantum numbers jp = 0+,predicted by sm are strongly favored [4, 5] ; • the pdg averaged mass is [6] : mh = 125.9 ± 0.4(stat) ± 0.4(syst) gev/c2 consistent with ew precision measurements that requires mh = 102−24−20 gev/c 2 [7]; • production cross section σpp→h agrees well inside the errors to the prediction of sm, that are affected by the uncertainty of 15%. • branching ratios to leptons, hadrons and gauge bosons are close enough to the sm predictions, with some tension in the channels h → γγ (µ ' 2) and h → bb (µ < 1) for atlas data. • angular distribution is slightly in favor of spin 1, but only with more data this could be confirmed, analyzing the decay h → ww → 2`2ν. at present, from february 2013 to november 2014, lhc is engaged in the first long shutdown aimed at the consolidation of the accelerator for running at the full design c.m.s. energy √ s = 14 tev. after this point, hopefully, supersymmetry (susy) hunting will be open. 2 supersymmetric higgs from the supersymmetry theory is expected a relation between the half-integer spin fermions to the integer spin bosons, introduced initially on a purely mathematical ground [8]. the real appeal of this theoretical framework is that it incorporates not only the three gauge fields of the sm but also gravity [9, 10]. figure 1: mssm model higgs masses, the hatched area indicates the decoupling susy parameters space ma > 2mz (adapted from ref. [11]). 42 http://dx.doi.org/10.14311/app.2014.01.0042 lhc, astrophysics and cosmology the particle spectrum of susy is extremely more crowded then the particle spectrum of sm. the present universe is constituted by matter fermions and force fields mediated by bosons, no susy particles has been identified. it means that at a time m−1pl ≤ tsusy ≤ tew the supersymmetry has been broken, in the sense that masses of susy partners became m̃susy � msm . for this reason the lightest neutral supersymmetric particles are excellent candidates for dark matter [12]. the higgs sector of susy theories is more complicated then in the sm. in the minimal supersymmetric extension of the sm, the higgs sector is schematically:( h0u h− ) ( h+ h0d ) a0 (1) namely a cp-odd pseudo scalar field a0 and two doublets of scalar cp-even fields. whose neutral components have v.e.v., assumed to be 〈 h0u 〉 = vu and〈 h0d 〉 = vd normalized to the value of fermi constant v2u + v 2 d = √ 2gf , with a ratio parameterized as vu/vd = tan β. the physical fields detectable at accelerator h and h are mixed states of the neutral components of the doublets with masses mh < mh, although it was never discarded the possibility to detect also the other states, especially the charged ones. if the mass of the pseudo-scalar ma and tan β are taken as free parameters, the masses of the other four are fixed by the equations [13]: m2h± = m 2 a + m 2 z + ∆m 2 h± m2h,h = 1 2 ( m2a + m 2 z ) + ∓ √ m2a + m 2 z − (2mamz cos 2β) 2 + ∆m2h,h (2) where ∆m2 h± and ∆m2h,h are the appropriated radiative corrections. fig. 1 shows the predictions of eq. (2) in which is clear that the light susy higgs mass saturates in the limit of decoupling ma � mz to mh → mz |cos 2β| + ∆m2h. the actual value of this limit, that depends strongly from the radiative corrections, is estimated to be in the range 130-150 gev/c2 [15, 16]. fig. (2) from cms [14] shows the result of the search of higgs like particles decaying to pairs of gauge bosons, giving evidence that a susy high mass higgs with mass lower than 700 gev/c2 is excluded. in facts the total production cross sections for the neutral susy higgs is essentially the same of that expected in sm, but the dominant decay channels are h → (w±,z) with vector bosons in final state [11]. fig. 2 shows the results of the cms search of higgs with mh > 200 gev/c 2 [14], that clearly excludes any mass mh ≤ 600 gev/c2 at 95% cl. figure 2: observed (solid line) and expected (dashed line) 95% cl upper limit on the ratio of the product of production cross section and branching ratio to the sm expectation for the higgs boson (from ref. [14]). 3 higgs boson in cosmology inflation is nowadays a well accepted paradigm in cosmology. the only problems is what is physically the “inflaton”. the original field considered as inflaton by guth was actually the higgs field [17], whose potential v (φ) = 1 4 λφ4 can naturally produce inflation if λ is small [18]. but this model predicts that the amplitude of density perturbation is δρ/ρ ∼ √ λ. therefore, in order to explain the observed δρ/ρ ∼ 10−4, it should be λ ∼ 10−10, absolutely irreconcilable with an higgs mass mh = √ 2λv, with v = (√ 2gf )−1 2 given by the strengths of the weak interactions. a solution could be a coupling of the higgs field with the gravity [19], with a lagrangian density lgr+sm = ( 1 2 m2pl + ξφ 2 ) r + 1 2 |∂µφ| 2 −vsm (φ) (3) where mpl = (8πgn ) −1 is the reduced planck mass, r the ricci scalar, vsm (φ) = 14λ ( φ2 −v2 )2 is the potential of the sm higgs field and ξ 6= 0 its coupling to the gravity. applying a rescaling of the metric g̃µν = ( 1 + 2ξφ2/m2pl )−1 gµν [20] the potential of the field becomes veff (φ) = λ 4 ( φ2 −v2 )2( 1 + 2ξφ 2 m2pl )2 (4) for the present small value of the higgs field φ � mpl/ √ 2ξ we have ṽ (φ) 'v(φ), while for φ & mpl/ √ 2ξ the effective potential of eq. (4), has a 43 giulio auriemma plateau that allows a successful slow-roll inflation. a first constraint on the coupling ξ is obtained from the amount of expansion of the universe reached at the end of the inflationary phase. the number of e-folding, defined from a(tf ) = a(ti) exp(−n) is predicted to be, using the slow-roll approximation [21], the higgs field that varies in the expansion from an initial value φi to a final one φf ≈ mpl/ √ 2ξ. the ratio is φi/φf ≈ √ 8n/3 for ξ � 1. the amplitude of anisotropy depends on the mass of the higgs boson, which fixes the value of the quartic coupling λ. the predicted scalar amplitude is δρ/ρ ≈ n π √ 18 √ λ/ξ. in order to justify the observed δρ/ρ ≈ 10−4 a strong coupling ξ & 104 is required. more refined calculations, including the non negligible radiative corrections to the higgs potential due the coupling with heavy quarks [22, 23], show that mh =126 gev/c 2 is compatible with the spectral index of the power law for scalar perturbations ns = 0.962 ± 0.002 and the upper limit for the tensor-to-scalar ratio r < 0.11 (95% c.l.) measured by planck [24]. theoretical studies have confirmed that susy inflation is possible, in many different scenarios [25] either with minimal [26] or non-minimal coupling [27]. as noted by linde [28] we can expect that due to the rich structure of the higgs sector in susy theories, the effective potential of the superfield will show several minima, interleaved by maxima where v ′(φ) � v (φ), suitable for slow-roll inflation in the minimal coupling case ξ = 0. the case of a non minimal coupling ξ > 0, is shown in ref. [27]. in a recent paper nakayama & takahashi [29] examined a susy model in which the lightest neutral higgs boson could be identified with the sm-like one. figure 3: planck best-fit value of the spectral index and mh = 126 gev/c 2 from lhc fits very well mssm inflaton model with small β (adapted from ref. [29]). 4 dark matter cosmological non-baryonic dark matter with ωdmh 2 = 0.120 ± 0.003(stat) ± 0.03(syst) has been estimated by planck [30]. figure 4: status of experimental direct search for wimps. the 68% cl region of possible positive detection are shown as filled area, while solid line represents 95% cl upper limits. the type of relics that can supply this density at “freeze out” should have a thermally averaged rate of annihilation at “freeze-out” (tf ' mdm/20) : 〈σav〉tf ' 2 × 10 −26 ( 0.12 ωdmh2 ) cm3s−1 (5) for mdm > 10 gev/c 2 [31]. it is worth noticing that the dm particles will be non-relativistic at freeze-out with β ≈ 4 × 10−2, that is dubbed “cold dark matter” (cdm). in addition the annihilation cross section that produces a rate comparable to the one of eq. (5) is close to the weak scale σ ≈ g2fm 2/16π for an hypothetical weakly interacting massive particle (wimp), not existing in the sm. the density of wimp’s in the solar neighborhood is estimated to be ρdmlocal = 0.39 ± 0.03 gev/cm 3 and is flowing with a velocity ≈ 200 km/s. wimp of mass 10 gev/c2, scattering against nucleons with a cross section σsi ≈ 10−40 cm2, give 0.5 interactions/100 kgday in a suitable detector[32]. the present controversial status of the direct searches is summarized in fig. 4. this figure shows that the claims for positive detections [33, 34, 35] seem to cluster around a “low” mass value mwimp ≈ 10 gev/c2 and a “high” elastic scattering cross section σsi ≈ 0.5 × 10−40 cm2/nucleon. particularly tantalizing is the claimed observation of a yearly modulation of the detected signal [36], that could be a clear signature of the association of the detected particles with the galactic halo. but unfortunately fig. 4 shows also that all but dama/libra experiments do not show evidence of a similar signal, even if it should be well inside their sensitivity [37]. 44 lhc, astrophysics and cosmology indirect searches for dm are based on the detection of radiation produced in the annihilation and decay of relic wimp’s. neutrino, gamma and antimatter astronomy are the basic tools of this search. from the eq. 5 it is possible to derive an order of magnitude of the annihilation rate. it is worth to stress however that this estimate is an average of σav over the thermal velocity distribution at tf � 1 gev, while annihilations in a galactic environment will take place at much lower temperature. moreover the source function of the astrophysical radiation will be qk = 〈σav〉bkρ2/m2 in photons cm−3s−1 being bk the inclusive branching ratio to the sm particle k, ρ and m respectively the local energy in gev/c2 cm−3 and the mass of the wimp in gev/c2. the spectrum of the γ-rays (or neutrinos) produced is composed by a continuum, extending up to the kynematical limit eγ . m and several monochromatic lines, each corresponding to a two body final states. the latter very attractive signature was proposed since 1988 by the compton gamma ray obsevatory [38]. in facts the first indication of a possible dm component in the diffuse galactic γ-rays from the galactic plane was given by the egret spark-chamber calorimeter [39], that found a significant excess on the galactic plane for eγ ≥ 1 gev. the poor energy resolution of the egret calorimeter for hard γ-rays did not allowed any search for lines, but the intensity and distribution in the galactic frame of this radiation was found to be close to what expected for annihilation or decay of particles with mass in the range 50-100 gev/c2. in the case of wimp’s the γ inclusive annihilation rate is σγv ≈ 10−26 cm3s−1 [40]. the diffuse γ−rays emission from selected regions on the galactic plane has been measured with high energy resolution (∼ 8% in the range 1-300 gev) by the csi scintillator trackercalorimeter of the fermi-lat instrument [41, 42]. the observed γ-rays spectrum is fitted with a single power law e−αγ for eγ ≥ 12.6 gev, with α = 2.44±0.01. narrow lines of width compatible with the instrumental resolution have been searched using background+signal max-likelihood method. fermilat has not detected any statistically significant γ-ray line in the range from 5 to 300 gev. an upper limit of the rates σ2γv ≤ 0.02 − 3.6 × 10−27 cm3s−1, for wimp’s masses 5 ≤ m ≤ 200 gev/c2, can be estimated from the flux upper limit at 95% c.l., applying the most optimistic galactic halo wimp density profile from n-body simulations [43]. if a simple isothermal profile is assumed, the limit increases by 50%. this limit starts to be close to the predicted rate for susy dm candidates [44]. unfortunately with this new data the claimed line at ≈130 gev [45, 46] corresponding to σ2γv ≈ (1.1 ± 0.5) × 10−27 cm3s−1 has not gained statistical significance (≈ 3.3σ). an excess of positrons in the energy range 4 ≤ ee ≤ 50 gev in cosmic rays, was discovered by a balloonborne instrument launched in 1974 from palestine, tx, in the heroic age of astroparticle physics [47]1. it is remarkable that after about 40 years we have a high precision mesurement of the positron flux covering the range 0.5 ≤ ee ≤ 350 gev from the 8.5 tons particle’s spectrometer ams-02 installed on the international space station [48]. the ams-2 positron fraction vs. energy is first decreasing from 8.42% at 1 gev to a minimum of 5.1% at 7 gev then increasing up to ' 15% at 260 gev. an excellent fit to this behaviour can be obtained either from pair emission by pulsar[49] or wimps annihilation [50]. in the latter case the mass of the wimp should be in the range 750 gev ≤ m ≤ 1.5 tev and inclusive annihilation rate 10−23 ≤ σe+v ≤ 10−22 cm3s−1[52] significantly larger then the one of eq. 5 ref. [51]. the discriminating observation is definitively the detection of an excess in the antiproton flux, because in the annihilation to lepton pair and quarks one is strongly correlated. up to now the pamela data [53] do not support any deviation from cosmic rays secondary production. alternatively, by assuming that the excess of positrons is all due to astrophysical sources, ref. [54] finds an upper limit to the annihilation rate that varies from 10−26 cm3s−1 at m = 10 gev/c2 to 10−23 cm3s−1 at m = 1 tev/c2, assuming that the dominant annihilation channel is µ+µ−. 5 lhc search for dark matter candidates the lower mass state of the four spin 1 2 neutralinos χ01,χ 0 2,χ 0 3,χ 0 4, predicted in r-parity conserving susy models, is a good candidate for the role of wimp. other candidates do exists, as for example the superpartner of the graviton, the spin 3 2 gravitino, that would be a superwimp because it couples with ordinary matter only via the gravitational interaction, making direct dm detection practically impossible [55]. at hadronic colliders such as lhc only wimps that couple with protons, such as the neutralino, can be directly produced. the production of neutralino candidate can be tested, quasi-model independently, using the process: pp → ( χ01χ̄ 0 1 ) + x (6) where x can be one (or more) hadronic jet, hard leptons or photons, while the neutralino pair (if r-parity is conserved) do not interact with the detector, but uppears in the kinematics of the event as missing trans1note that the young member of this slac team will be the nobel laureate of 2006 for cobe. 45 giulio auriemma verse energy emisst [56]. the dominant sm physical background for the reaction (6) is pp → w±/z0 → ``, where the leptons are neutrinos or are not detected, that can be subtracted and/or reduced by optimized kinematical cuts. figure 5: upper limits on σsi (left) and σ2γv (right) inferred from monojet+emisst atlas data [57]. considering the simpler case that x is a single hadronic jet the cross section for neutralino pairs production, that contributes to process (6), will be σqq→χχ̄ ≈ αsg2χg2qp jet t /m 4 ∗ where αs = 0.64 is the qcd constant, gχ,gq the susy couplings being q = q, q̄ or g and finally m∗ � pt is suppressing mass scale. the detection of a number of events significantly larger the the calculated sm background would be strong indication of a wimp candidate, whose mass could be inferred from the missing energy distribution. in addition a comparison with direct dm search experiments, discussed in the previous section, would be possible because the scattering cross section, from the same coupling, is predicted to be σsi ≈ g2χg2qµ2/m4∗ being µ the reduced mass of the wimp-nucleon system [58]. moreover the same suppression factor m−4∗ enters, together with phase space factor depending only from the masses, in the annihilation rate of the neutralino (see for example eq. (10) and (11) of ref. [59]) that allows comparison with the dm relic density and the indirect searches. fig. (5) shows the potentiality of this type of searches at lhc [57, 60]. figure 6: lhcb signal for b0s → µ+µ− [62] (left) and presently cmssm allowed region (green) on the plane (m0,m1/2) [55] (right). important constraints on the supersymmetries, beyond the effective field theories discussed before, that can also set limits to the wimp candidates properties, are given by the search for rare decays of heavy neutral mesons like the b0d and b 0 s decays, copiously produced at lhc [61], performed by the lhcb experiment as well as the cms experiment. lhcb is the lhc experiment devoted to flavor physics. both lhcb and cms have measured the br(bs → µµ) which is an important test bench for new physics, because it strongly suppressed in the minimal smit is enhanced by susy. limit at 95% cl is br(bs → µµ)susy < br(bs → µµ)sm . 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[63] d. bauer, et al. in ”report prepared for the community summer study”, (snowmass) 2013 arxiv: 1305.1605. 48 http://arxiv.org/abs/1007.0821 http://arxiv.org/abs/1309.2570 http://dx.doi.org/10.1103/physrevlett.105.121101 http://arxiv.org/abs/1304.1414 http://arxiv.org/abs/1008.1783 http://arxiv.org/abs/1210.4491 http://dx.doi.org/10.1103/physrevd.82.116010 http://arxiv.org/abs/1005.3797 http://arxiv.org/abs/1109.4398 http://arxiv.org/abs/1207.7315 http://arxiv.org/abs/1307.5024 http://arxiv.org/abs/1307.5024 http://arxiv.org/abs/1305.1605 http://arxiv.org/abs/1305.1605 introduction supersymmetric higgs higgs boson in cosmology dark matter lhc search for dark matter candidates summary 116 acta polytechnica ctu proceedings 2(1): 116–122, 2015 116 doi: 10.14311/app.2015.02.0116 inner disk structure of dwarf novae in the light of x-ray observations ş. balman1 1middle east technical university, department of physics, inonu bulvarı, ankara, 06531, turkey corresponding author: solen@astra.physics.metu.edu.tr abstract diversity of the x-ray observations of dwarf nova are still not fully understood. i review the x-ray spectral characteristics of dwarf novae during the quiescence in general explained by cooling flow models and the outburst spectra that show hard x-ray emission dominantly with few sources that reveal soft x-ray/euv blackbody emission. the nature of aperiodic time variability of brightness of dwarf novae shows band limited noise, which can be adequately described in the framework of the model of propagating fluctuations. the frequency of the break (1-6 mhz) indicates inner disk truncation of the optically thick disk with a range of radii (3.0-10.0)×109 cm. the rxte and optical (rtt150) data of ss cyg in outburst and quiescence reveal that the inner disk radius moves towards the white dwarf and receeds as the outburst declines to quiescence. a preliminary analysis of su uma indicates a similar behaviour. in addition, i find that the outburst spectra of wz sge shows two component spectrum of only hard x-ray emission, one of which may be fitted with a power law suggesting thermal comptonization occuring in the system. cross-correlations between the simultaneous uv and x-ray light curves (xmm −newton) of five dne in quiescence show time lags in the x-rays of 96-181 sec consistent with travel time of matter from a truncated inner disk to the white dwarf surface. all this suggests that dwarf novae and other plausible nonmagnetic systems have truncated accretion disks indicating that the disks may be partially evaporated and the accretion may occur through hot (coronal) flows in the disk. keywords: cataclysmic variables white dwarfs accretion, accretion dics stars: binaries x-rays: stars radiation mechanisms: thermal stars: dwarf novae. 1 introduction dwarf novae (dne) are a class of cataclysmic variables (cvs) which are interacting compact binaries in which a white dwarf (wd, the primary star) accretes matter and angular momentum from a main (or post-main) sequence star (the secondary) filling its roche-lobe. the matter is transferred by means of an accretion disk that is assumed to reach all the way to the wd surface. ongoing accretion at a low rate (quiescence) is interrupted every few weeks to months or sometimes with longer durations by intense accretion (outburst) of days to weeks where ṁ increases (see warner 1995 for a review). the material in the inner disk of nonmagnetic cvs initially moving with keplerian velocity dissipates its kinetic energy in order to accrete onto the slowly rotating wd creating a boundary layer (bl) (see mauche 1997, kuulkers et al. 2006 for an overview). standard accretion disk theory predicts half of the accretion luminosity to originate from the disk in the optical and ultraviolet (uv) wavelengths and the other half to emerge from the boundary layer as x-ray and/or extreme uv (euv)/soft x-ray emission which may be summerized as lbl∼ldisk=gmwdṁacc/2rwd=lacc/2 (lyndenbell & pringle 1974, godon et al. 1995). during low-mass accretion rates, ṁacc<10 −(9−9.5)m�, the boundary layer is optically thin (narayan & popham 1993, popham 1999) emitting mostly in the hard xrays (kt∼10(7.5−8.5) k). for higher accretion rates , ṁacc≥10−(9−9.5)m�, the boundary layer is expected to be opticallly thick (popham & narayan 1995) emitting in the soft x-rays or euv (kt∼10(5−5.6) k). the transition between an optically thin and an optically thick boundary layer, also depends on the mass of the white dwarf (also rotation) and on the alpha viscosity parameter. 2 x-ray observations of dwarf novae in quiescence and outburst the quiescent x-ray spectra are mainly characterised with a multi-temperature isobaric cooling flow model of plasma emission at tmax=6-55 kev with accretion rates of 10−12-10−10 m� yr −1. the x-ray line spectroscopy indicates narrow emission lines (brightest oviii kα) and near solar abundances, with a 6.4 kev line expected to be due to reflection from the surface of the wd. the detected doppler broadening in lines during quies116 http://dx.doi.org/10.14311/app.2015.02.0116 inner disk structure of dwarf novae in the light of x-ray observations cence is <750 km s−1 at sub-keplerian velocities in the boundary layer with electron densities >1012 cm−3 (see baskill et al. 2005, kuulkers et al. 2006, rana et al. 2006, pandel et al. 2005, balman et al. 2011, balman 2012). the total x-ray luminosity during quiescence is 1029 -1032 erg s−1. lack of bl emission in the xrays have been suggested due to low lx/ldisk ratio (see kuulkers et al. 2006). it has been suggested for dn in quiescence that if the wd emission is removed and some disk truncation is allowed, this ratio is ∼1 (pandel et al. 2005). dn outbursts are brightenings of the accretion disks as a result of thermal-viscous instabilities summerized in the dim model (disc instability model; lasota 2001,2004). during the outburst stage, dn x-ray spectra differ from the quiescence since the accretion rates are higher (about two orders of magnitude), the bl is expected to be optically thick emitting euv/soft xrays (lasota 2001, see the x-ray review in kuulkers et al. 2006). on the other hand, soft x-ray/euv temperatures in a range 5-25 ev are detected from only about five systems (e.g., mauche et al. 1995, mauche & raymond 2000, long et al. 1996, mauche 2004, byckling et al. 2009). as a second and more dominantly detected emission component, dn show hard x-ray emission during the outburst stage however, at a lower flux level and x-ray temperature compared with the quiescence (e.g., ww cet & su uma: collins & wheatley 2010, fertig et al. 2011, ss cyg: wheatley et al. 2003, macgowen et al. 2004, ishida et al. 2009). on the other hand, some dn show increased level of x-ray emission (gw lib & u gem: byckling et al. 2009, guver et al. 2006). the total x-ray luminosity during outburst is 1030 -1034 erg s−1. the grating spectroscopy of the outburst data indicate large widths for lines with velocities in excess of 1000 km s−1 mostly of h and he-like emission lines (c,n,o,ne, mg, si, fe, ect.) (mauche 2004, rana et al. 2006, okada et al. 2008). a characteristic of some dn outburst light curves are the uv and x-ray delays in rise to outburst (w.r.t. optical) indicating possible disk truncation ((meyer & meyer-hofmeister 1994 and references therein) . these delays are a matter of several hours (upto a day) that need dedicated simultaneous multi-wavelength observations. during outburst no eclipses are detected in the eclipsing systems (particularly of soft x-ray emission) or no distinct orbital variations are seen. (e.g., pratt et al. 1999, byckling et al. 2009). note that both in ss cyg (mc gowan et al. 2004) and in su uma (collins & wheatley 2010) the x-ray flux in between outbursts have been found to decrease as opposed to expectations of the dim model. wz sge is a short period su uma type dwarf nova with long interoutburst interval of 20-30 yrs. the system is believed to show no normal outbursts but only superoutbursts. the most recent outburst in julyaugust 2001 (previous outbursts in 1978, 1946 and 1913) was observed from ground and space over several different wavelength bands including the x-ray wavelengths and the euv using the chandra observatory (see wheatley & mauche 2005, kuulkers et al. 2002). there has been three letg (low energy transmission grating) and three acis-s (advanced camera imaging specreometer) observations obtained in the continuousclocking (cc) mode. i report a preliminary analysis on the spectral fitting of the acis-s cc-mode data. figure 1: chandra x-ray outburst spectra of wz sge in 0.3-10.0 kev range obtained with acis-s (ccmode). the observation days after the outburst are labeled on each spectrum. the spectra and response files were prepared in a standard way (task specextract is used) by choosing the source and the background photons using a suitable extraction box (10 arcsec in size) on the data strip using ciao 4.4 and the caldb 4.4.8 (see http://cxc.harvard.edu/ciao/). for further analysis, heasoft version 6.13 is utilized. the spectra were fitted with the multi-temperature plasma emission models (e.g., cevmkl in xspec, see http://heasarc.nasa.gov/xanadu/xspec/) according to the expections from earlier work of dn analysis in quiescence and outburst. cevmkl model is a multitemperature plasma emission model (built from mekal code, mewe et al. 1986) where emission measures follow a power-law in temperature (dem = (t/tmax)α−1 dt/tmax). the reduced χ2 of the fits were above a value of 2 with either a single mekal model or cevmekl alone (the α parameter of the power law distribution of temperatures is set free) for days 6, 15, 30 after outburst. i added a second component of a power law which reduced the χ2 values to desirable levels. in general, the plasma has variable abundance of n,o,ne,fe, and s. the spectral parameters for day 6 after the outburst were ktmax=0.7-1.3 kev with a photon index of γ=0.8-1.4. the x-ray flux of the ther117 ş. balman mal cevmkl component was 1.1×10−11 erg s−1 cm−2 and lcevmkl is 2.7×1030 erg s−1 where the power law component has 5.8×10−12 erg s−1 cm−2 and lpower is 1.3×1030 erg s−1 (43.5 pc is assumed, harrison et al. 2004). the parameters for day 15 were ktmax=0.5-1.5 kev and γ=0.2-1.1. between these two dates the x-ray flux of the cevmkl component decreased by 2 and the power law component inceased by about 1.4 . for day 30 ktmax=1.2-3.0 kev and γ=1.5-2.0. therefore, the x-ray emitting region gets hotter and the photon index decrease. the x-ray fluxes recover to day 6. the final observation on day 58 which is almost quiescence after the outburst shows ktmax=26-46 kev and no power law emission is detected. the x-ray flux is the highest 4.4×10−11 erg s−1 cm−2 and lx is 1.1×1031 erg s−1 (at 43.5 pc). all ranges correspond to 90% confidence level errors. the neutral hydrogen absorption has stayed at the interstellar level in all fits within errors. no optically thick blackbody emission is detected and only x-ray suppression of the quiescent emission is observed just like some other dn. the x-ray temperatures on days 6-30 are in good agreement with the letg results as well. a model of three additive mekals yield acceptable fits on days 6-30, however allowing for very high x-ray tempartures. a partial covering absorber model improves the cevmkl fits, however the high intrinsic absorption when modeled yields inconsistent count rates for the letg observations. 3 inner disk structure using eclipse mapping techniques flickering is a fast intrinsic brightness scintilation occurring on time scales from seconds to minutes (e.g., amplitudes of 0.01 1 mag in the optical). it is observed in all accreting sources. eclipse mapping methods have been used to reproduce the spatial distribution of flickering in cvs in the optical and uv wavelengths. some eclipse mapping studies of flickering in quiescent dwarf novae indicate that mass accretion rate diminishes by a factor of 10-100 and sometimes by 1000 in the inner regions of the accretion disks as revealed by the brightness temperature calculations which do not find the expected r−3/4 radial dependence of brightness temperature in standard steady-state constant mass accretion rate disks (see e.g., z cha and oy car: wood 1990, v2051 oph: baptista & bortoletto 2004, v4140 sgr: borges & baptista 2005). biro (2000) finds that this flattening in the brightness temperature profiles may be lifted by introducing disk truncation in the quiescent state (e.g., r ∼ 0.15rl1 ∼4×109 cm; dw uma, a nova-like). a comprehensive uv modeling of accretion disks at high accretion rates in 33 cvs including many nova-likes and old novae (puebla et al. 2007) indicate an extra component from an extended optically thin region (e.g., wind, corona/chromosphere) evident from the strong emission lines and the p cygni profiles. this study also indicates that the mass accretion rate may be decreasing 1-3 orders of magnitude in the inner disk region (divergence starts r≤ afew ×1010 cm towards the wd). 4 inner disk structure of dwarf novae the truncation of the optically thick accretion disk in dne in quiescence was invoked as a possible explanation for the time lags between the optical and uv fluxes in the rise phase of the outbursts (meyer & meyerhofmeister 1994), and for some implications of the dim (see lasota 2001, 2004) or due to the unusual shape of the optical spectra or light curves of dne (linell et al. 2005, kuulkers et al. 2011). theoretical support for such a two-phase flow was given by a model of the disk evaporation of (meyer & meyer-hofmeister 1994). this model was later elaborated to show that the disk evaporation (coronal “syphon” flow) may create optically thick-optically thin transition regions at various distances from the wd (liu et al. 1997, mineshige et al. 1998). 4.1 propagating fluctuations model another diagnostic tool proposed to study the inner disk structure in accreting objects is the aperiodic variability of brightness of sources in the x-rays. while the long time-scale variability might be created in the outer parts of the accretion disk (warner & nather 1971), the relatively fast time variability (at f >few mhz) originates in the inner parts of the accretion flow (bruch 2000; baptista & borteletto 2004). properties of this noise is similar to that of the x-ray binaries with neutron stars and black holes. now, the widely accepted model of origin for this aperiodic flicker noise is a model of propagating fluctuations (lyubarskii 1997, churazov et al. 2001, uttley & mchardy 2001, revnivtsev et al. 2009,2010, uttlet et al. 2011, scaringi et al. 2012).the modulations of the light are created by variations in the instantaneous value of the mass accretion rate in the region of the energy release. these variations in the mass accretion rate, in turn, are inserted into the flow at all keplerian radii of the accretion disk due to the stochastic nature of its viscosity and then transferred toward the compact object. thus, variations are on dynamical timescales. this model predicts that the truncated accretion disk should lack some part of its variability at high fourier frequencies, i.e. on the time scales shorter than a typical time scale of variability. 118 inner disk structure of dwarf novae in the light of x-ray observations 4.2 broad-band noise of dwarf novae a recent work by balman & revnivtsev (2012) have used the broad-band noise characteristic of selected dn in quiescence (only one in outburst: ss cyg) and studied the inner disk structure and disk truncation via propagating fluctuations model. the power spectral densities (pds) expressed were calculated in terms of the fractional rms amplitude squared following from (miyamoto et al. 1991). the light curves were divided into segments using 1-5 sec binning in time and several pds were averaged to create a final pds of the sources. the white noise levels were subtracted hence leaving us with the rms fractional variability of the time series in units of (rms/mean)2/hz. next, the rms fractional variability per hertz was multiplied with the frequencies to yield νpν versus ν. for the model fitting a simple analytical model is used p(ν) ∝ ν−1 ( 1 + ( ν ν0 )4)−1/4 , which was proposed to describe the power spectra of sources with truncated accretion disks (see revnivtsev et al. 2010,2011). the broad-band noise structure of the keplerian disks often show ∝ f−1...−1.3 dependence on frequency (churazov et al. 2001, gilfanov et al. 2005), and this noise will show a break if the optically thick disk truncates as the keplerian motion subsides. balman & revnivtsev (2012) show that for five dn systems, ss cyg, vw hyi, ru peg, ww cet and t leo, the uv and x-ray power spectra show breaks in the variability with break frequencies in a range 1-6 mhz, indicating inner disk truncation in these systems. the truncation radii for dn are calculated in a range ∼(310)×109 cm including errors (see table 2 in balman & revnivtsev 2012). the same authors used the archival rxte data of ss cyg in quiescence and outburst listed in table 1 of their paper to derive the broad-band noise of the source in different states (i.e., accretion rates). for the outburst phases, authors investigated times during the x-ray suppression (e.g., the x-ray dips; optical peak phases of the outburst) and the x-ray peak. this, in turn is expected to indicate the motion of the flow in the inner regions of the disk and the geometry of the inner disk. authors show that the disk moves towards the white dwarf during the optical peak to ∼ 1×109 cm and receeds as the outburst declines to quiescence to 5-6×109 cm. this is shown for a cv, observationally, for the first time in the x-rays (see figure 2). this is also supported with the optical data analysis of ss cyg in quiescence and outburst (revnivtsev et al. 2012) 4.3 x-ray and uv light curve cross-correlations balman & revnivtsev (2012) calculated the crosscorrelation between the two simultaneous light curves (x-ray and uv), using the archival xmm-newton epic pn and om data utilizing heasoft task crosscor. to obtain the ccfs (cross-correlation functions), datasets were divided into several pieces using 1-5 sec binning in the light curves and fitted the resulting ccfs with double lorentzians since it was necessary for adequate fitting (see the paper for details). the ccfs for all the dne show clear asymmetry indicating that some part of the uv flux is leading the x-ray flux. in addition, they detect a strong peak near zero time lag for ru peg, ww cet and t leo, suggesting a significant zero-lag correlation between the x-rays and the uv light curves. figure 2: the rxte pds of ss cyg in quiescence and during the optical peak of outburst (top). the solid lines show the fit with the propagating fluctuations model along with two lorentzians. the same authors also calculated the crosscorrelation between the simultaneous uv and x-ray light curves by subtracting the zero time lag components in the five dne pds, yielding more correct time lags consistent with delays in the x-rays of 96-181 sec (see the paper on details of the modeling). the lags occur such that the uv variations lead x-ray variations which shows that as the accreting material travels onto the wd, the variations are carried from the uv into the x-ray emitting region. therefore, the long time lags of the order of minutes can be explained by the travel time of matter (viscous flow) from a truncated inner disk to the white dwarf surface. zero time lags (∼ light travel time) indicate irradiation effects in these systems since we do not have resolution better than 1 sec. 119 ş. balman figure 3: see figure 10 in balman & revnivtsev (2012). the cross-correlation of the epic pn (x-ray) and om (uv) light curves with 1 sec time resolution. the two-component lorenzian fits are shown as solid black lines (except for ss cyg). figure 4: the rxte pds of su uma in quiescence and in optical outburst is displayed (top panel). xmmnewton epic pn pds of v426 oph and ht cas in quiescence (middle and bottom panel). the solid lines are the fits with the propagating fluctuations model. an analysis on the rxte data of su uma in quiescence and outburst following six consecutive outburts reveal a similar broad-band noise structure to ss cyg in quiescence showing a break frequency ∼ 5.5-7.5 mhz with a truncated optically thick disk ∼ 3.8×109 cm. the preliminary analysis of the outburst data during the x-ray suppression episodes (optical peak of the outburst) indicates no disk truncation or a truncation around 0.1 hz (see figure 4). therefore, su uma indicates a similar behaviour of the disk during the outburst to ss cyg where the inner disk moves in almost all the way to the wd during the optical peak and moves out in decline to quiescent location further out. the lower level of broadband noise during outburst stage of su uma may be due to the high radiation pressure support of an optically thick disk flow during the peak of the outburst suppressing the variations. a similar pds analysis of the xmm-newton data of the dne, v426 oph and ht cas in quiescence reveals break frequencies 4.6-2.6 mhz and 7.2-2.8 mhz, respectively. the approximate keplerian radii where the optically thick disk truncates is ∼ 6.2×109 cm and ∼ 4.5×109 cm, respectively (see figure 4). 5 conclusions studies of dne broad-band noise characteristics in the x-rays (see also balman & revnivtsev 2012) indicate that dne have truncated accretion disks in quiescence detected in at least 8 systems in a range ∼(3.0-10.0)×109 cm including errors. the magnetic cvs (mcvs) show rather smaller truncation radii (0.9120 inner disk structure of dwarf novae in the light of x-ray observations 2.0)×109 cm (revnivtsev et al. 2010, 2011). this can also explain the uv and x-ray delays in the outburst stage and the accretion may occur through hot (coronal) flows in the disk. note that extended emission and winds are detected from dn in the outburst stage which may be an indication of the coronae/hot flows in these systems (e.g., mauche 2004). time delays detected in a range of 96-181 sec, are also consistent with matter propogation timescales onto the wd in a truncated nonmagnetic cv disk in quiescence. balman & revnivtsev (2012) approximate an α of 0.1-0.3 for the inner regions of the dne accretion disks in quiescence. in addition, the outburst spectra of wz sge shows two component spectrum of only hard x-ray emission, one of which may be fitted with a power law suggesting thermal comptonization of the optically thick disk photons or scattering from an existing wind during the outburst. the spectral evolution and disapearance of the power law component after outburst may support this issue and that the accretion disk goes back to its quiescent truncated structure and comptonization stops. a general picture of the accretion flow around a wd in quiescence thus might be somewhat similar to that of the black hole/neutron star accretors with an optically thick colder outer accretion disk and an optically thin hot flow in the inner regions where the truncation occurs (see esin et al. 1997, done et al. 2007). the appearance of a hot flow (e.g., adaf-like) in the innermost regions of the accretion disk will differ from that of ordinary rotating keplerian disk because it is no longer fully supported by rotation, but might have a significant radial velocity component with sub-keplerian speeds. it is important to monitor dne in the x-rays to measure the variability in the light curves in time 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[58] wilms, j., allen, a., & mccray, r. 2000, apj, 542, 914 [59] wood, j. h.: 1990, mnras, 243, 219 122 http://dx.doi.org/10.1086/421438 http://dx.doi.org/10.1086/309489 http://dx.doi.org/10.1086/380758 http://dx.doi.org/10.1038/362820a0 http://dx.doi.org/10.1086/587161 http://dx.doi.org/10.1086/429983 http://dx.doi.org/10.1111/j.1365-2966.2012.20512.x http://dx.doi.org/10.1046/j.1365-8711.2001.04496.x http://dx.doi.org/10.1111/j.1745-3933.2011.01056.x http://dx.doi.org/10.1017/cbo9780511586491 http://dx.doi.org/10.1093/mnras/152.2.219 http://dx.doi.org/10.1111/j.1365-2966.2003.07149.x http://dx.doi.org/10.1086/317016 introduction x-ray observations of dwarf novae in quiescence and outburst inner disk structure using eclipse mapping techniques inner disk structure of dwarf novae propagating fluctuations model broad-band noise of dwarf novae x-ray and uv light curve cross-correlations conclusions 189 acta polytechnica ctu proceedings 1(1): 189–193, 2014 189 doi: 10.14311/app.2014.01.0189 are supernovae responsible for the gamma ray spectrum from the galactic center? todor stanev1 1bartol research foundation and department of physics and astronomy, university of delaware, u.s.a. corresponding author: stanev@bartol.udel.edu abstract we discuss the supernova remnants distribution as a function of the galactic longitude and compare their positions to that of the detected tev gamma ray sources. a large fraction of these sources either coincide or a close by known supernova remnants. if we look within 10o of the galactic center most identified sources are combinations of supernova remnants with giant molecular clouds. the diffuse gamma ray flux from the direction of the galactic center is much smaller. keywords: supernova remnants cosmic rays acceleration tev gamma rays. 1 introduction i will start this write-up the same way as i started the talk: with the answer to the question, which is i do not know. what i will write about is what we know about different types of supernova remnants (snr) from which tev γ rays are detected and what theoretical research says about the way these γ rays are produced in them. the main reason we are interested in supernova remnants is they are supposed to be the sources of galactic cosmic rays, which are accelerated at the supernova shocks. the outer regions of the expanding supernovae initially move with high velocity, approximately 0.1c. the expansion velocity is supersonic and a shock is formed in front of the remnant. the shock collects the interstellar matter that it meets during the expansion. once the shock radius is close to 1 pc, when the remnant is about 1,000 years old, the amount of the swept-up material becomes too much and the remnants velocity decreases. this is the time when cosmic rays are accelerated at the shock front. the magnetic field at the shock front is significantly higher than that of the interstellar medium and only a small fraction o(few %) of the kinetic energy of the remnant can supply all galactic cosmic rays (ginzburg & syrovatskii, 1964). 2 supernova remnants the latest supernova remnant catalog that i am familiar with is that of d.a. green (green 2009). it contains 274 supernova remnants studied in radio and gives their location, power at 1 ghz, and spectral indexes. ten of these snr have longitude less than 10o from the galactic center. all of them are closer than 1.5o from the galactic plane. the directions of all 274 these supernova remnants are shown in fig. 1 where the ones from supernovae after 1,000 ad (and the galactic center) are indicated. one can easily see that most of the remnants are very close to the galactic plane and a few are more than 5o away from it. one can also see that the supernova remnants density is much higher in the inner 60o of the galaxy. figure 1: locations of the supernova remnants in d.a. green catalog. the question if these snr are the sources of cosmic rays and of high energy γ-rays becomes very reasonable. to approach this question we will have a look at the galactic sources of high energy γ-rays, the tevcat of the university of chicago (http://tevcat.uchicago.edu). this catalog contains 145 gamma ray sources that have been discovered by the atmospheric cherenkov gamma ray telescopes. these devices consist of several mirror telescopes that observe the cherenkov light emitted by the cascade developing after the high energy gamma rays interact with atmosphere. the shape of the image of the cascade is enough to differentiate between atmospheric γ-ray cascades and cascades from the interactions of cosmic ray 189 http://dx.doi.org/10.14311/app.2014.01.0189 todor stanev protons and nuclei in the atmosphere. when the cascade is observed by more than one telescope the angular resolution is a small fraction of a degree. the threshold energy for these devices depends on the size of the mirrors and is often of o(100 gev). the three major tev gamma ray atmospheric telescopes are, in order of their completion, hess (hinton et al, 2004), veritas (hanna et al., 2008), and magic (boria tridon et al., 2010). hess and veritas have each four 12 m. telescopes and magic has two 17 m. telescopes. the average detection threshold for 12 m. telescopes is between 100 and 200 gev while the 17 m. telescopes can come down to 60 gev. tevcat has identified 64 galactic sources listed by different types in table 1. table 1: galactic tev gamma ray sources pulsar wind nebulae 31 shell supernova remnants 12 x-ray binaries 3 gamma ray binaries 1 massive stars 3 snr/molecular clouds 8 globular clusters 1 unidentified 5 it is not obvious that the different distribution of the tev gamma ray sources within 10o of the galactic center is important. out of the ten sources four are snr with close by molecular cloud, three are unidentified (which includes the galactic center and the galactic center ridge), one is pulsar wind nebula. there is also one shell snr and one globular cluster. since all of these objects are at the approximately the same distance from us we expect the gamma ray fluxes from them to correspond to the overall luminosity of the sources. the number of the detected gamma rays will also correspond to the threshold energy for detection. the exact threshold energy for detection by the tev air cherenkov telescopes depends on the size of the telescope mirrors and on the positions of the source and the telescope. to demonstrate this we show in fig. 2 the galactic plane and the sources in equatorial coordinates. the telescopes are sensitive to source elevation down to 30o but the lower the source is the higher the energy threshold. the hess telescope is in namibia at 23 degrees south, veritas is in arizona, u.s.a. at 31 degrees north, and magic is at 29 degrees north. all of them can observe most of the galactic plane, but the location of hess is the best one for observations of the galactic center. this is the reason it was constructed in namibia. figure 2: the vicinity of the galactic plane and the galactic sources of tev gamma rays in equatorial coordinates. figure 3 shows the galactic sources of tev gamma rays (circles) overlayed on the supernova remnants. the sources from the first fermi/lat catalog (abdo et al (2010)) are shown with triangles. if we had a look at the later fermi/lat catalogs we would find many more galactic sources the coincide with supernova remnants. one should not forget that fermi/lat is sensitive to γrays of energy 0.1 to 100 gev and the average spectral slope of the gamma ray sources is about 2.5. this means that are 30,000 more gamma rays above 100 mev than are gamma rays above 100 gev. this number varies, of course, from source to source depending on the spectral index of the source. figure 3: the position of the galactic tev gamma ray sources (circles) are overlayed on top of the supernova remnants from the green’s catalog. the triangles show the the fermi/lat gamma ray sources from its first catalog (abdo et al (2010)). 2.1 pulsar wind nebulae pulsar wind nebulae are formed by the highly relativistic mhd winds expelled by the rotating neutron stars. such objects can accelerate all kinds of charged particles, from electrons to heavy nuclei if there is a proper injection mechanism close to the neutron star. the best studied pwn is that of the crab. the gamma ray fluxes from different sources are often given in crab units that describe the ratio of their emission to that of the crab. its gamma ray energy spectrum is best described by the synchrotron self compton model that is based only on electron acceleration. electrons emitted by the source suffer from synchrotron energy loss to photons. these photons than go through inverse compton interactions 190 are supernovae responsible for the gamma ray spectrum from the galactic center? with the accelerated electrons that pushes them to tev energies. it is not expected that such sources can produce very high energy gamma rays when the inverse compton cross section decreases in the klein-nishina regime. the studies of most tev pulsar wind nebulae, however, agree better with leptonic models than with cosmic ray interaction models. the same is true for many other tev gamma ray sources such as geminga and vela x. 2.2 shell supernova remnants shell supernova remnants sources are the typical supernova remnants as we imagine them. the shock wave from supernova explosion heats up the interstellar material as it propagates through it. when we observe such a remnants we see mostly its outer edge that is brighter than the inside of the remnant. in optical astronomy this effect is called limb brightening. the observations also indicate higher magnetic fields at the limbs of the remnants. a very good introduction to the processes in the remnant is given in (reynolds, 1998) where the author deals only with electron acceleration. the acceleration of protons and heavier nuclei is discussed by (caprioli, amato & blasi, 2010). in order to understand the limb brightening one has to assume that there is electron acceleration at the edge of the remnants where the electrons have synchrotron energy loss in the higher magnetic field. the shell supernova tev gamma ray sources in tevcat include ic443, sn1006, cassiopeiaã, and tycho. 2.3 snr/molecular clouds this type of sources includes supernova remnants that have massive molecular clouds nearby. very often the tev gamma ray sources do not point at the center of the remnant, rather at the molecular cloud itself or at the side of the remnant that is close to the cloud. figure 4: positions of the molecular clouds and the number of tev gamma rays observed by the hess experiment (aharonian et al, 2006). most of these sources were first observed by the hess experiment and carry its name. an extremely interesting analysis was performed by hess after its observation of the galactic center ridge (aharonian et al, 2006). this analysis is illustrated in fig. 4. hess saw that the number of tev gamma rays coming from the vicinity of the galactic center has gaps, i.e. there would be hundreds of gamma rays coming from certain direction, then almost no gamma rays, then again hundreds of gamma rays. the exact directions of the locations that produced hundreds of gamma rays coincided with the positions of several giant molecular clouds containing 2-4×107 m�. at the 8.5 kpc distance of the galactic center 0.2o longitudinal difference corresponds to a distance of 30 pc. the analysis concluded with the idea that cosmic rays were accelerated in the galactic center (maybe at sagittarius a east) many years ago and started diffusing away from it. when they diffuse into one of the huge molecular clouds they interact with the matter there, generate neutral mesons that decay into gamma rays. these cosmic rays were able to diffuse at distances up to 100 pc but have not yet reached the molecular cloud at a distance of 200 pc. this leads to an estimate of the time of the supernova remnant acceleration episode of 104 years ago. an acceleration episode could be similar to the movement of a small molecular cloud to the black hole sgr a∗ in the galactic center that is being observed now (gillesen et al, 2013). the absorption of large amount of matter by the black hole can easily create a shock and thus accelerate cosmic rays for a relatively short time. there are now analyses of different gamma ray sources that require the existence of molecular clouds nearby. a very interesting one is that of (torres, rodriguez marrero & decea del pozo, 2010) of the supernova remnant ic443. the gamma ray emission of this source is known from the 1990’s when it was detected by egret. it was observed by the tev cherenkov telescopes magic and veritas at a slightly different (0.4o) direction. the direction of the same source from agile and fermi/lat are consistent with this of egret. the authors of this analysis identify two (or maybe three) different sources: ic443 as seen by the tev telescopes and a molecular cloud in front of it that is observed in the gev energy range. it is also possible that another, relatively small molecular cloud also emits gamma rays. the existence of different sources solves the problem with the different spectral indices in the gev and the tev gamma ray emission. 3 other possible ideas an old paper (berezinsky et al, 1993) looks at the possibility that there would be a strong diffuse radiation from the central region of the galaxy. the paper uses a matter density study (bloemen, 1989) that determined that 191 todor stanev the matter density in the inner 300 pc of the galactic plane within b < |2|o is 38 nucleons cm−3. the matter density strongly decreases to reach less than 1 per cm3 at distances greater than 6 kpc. assuming that the matter distribution in the galaxy is symmetric (which it is not) the paper used the data of bloemen to calculate the column density in different directions. the highest column density for latitude less than 2o around the galactic center reached 8×1022 cm−2. using this mapping the paper provides the ratio between the γray flux and cosmic ray flux as a function of the galactic longitude. in the vicinity of the galactic center and for energies less than 10 tev this ratio is still less than 10−4. it is easy the explain that: a column density of 8×1022 cm−2 is slightly more than 0.1 g/cm2 when the proton interaction length is about 50 g/cm2. this means that only 0.6% of the galactic cosmic rays interact and generate gamma rays that we will see coming from the direction of the galactic center. in all other directions the diffuse gamma ray flux is smaller. i refer to this paper because it attempted to calculate in an easy way the diffuse gamma ray flux. the contemporary attempts to do this are much more sophisticated and involve measurements and subtraction of all known gamma ray sources. the fermi bubbles were discovered in this way (su, slatyer & finkbeiner, 2010). the bubbles cover an area much larger than the galactic center region and are still discussions of the origin of the gamma ray emission from them. one example of the fermi/lat studies of the diffuse gamma ray radiation in the galaxy is in (ackermann et al, 2012) where the small and large scale anisotropy is studied and the existence of unknown gamma rays sources is discussed. in other papers the diffuse radiation is searched for for possible dark matter signatures. 4 discussion and conclusions the gamma ray energy spectra from the detected gamma ray sources are very often compared to leptonic and hadronic models models of gamma ray production. some of the sources, typically the pulsar wind nebulae, fit better the leptonic models, where electrons accelerated at the source have inverse compton collisions with synchrotron photons emitted by the same electrons in propagation of the magnetic fields around the source. even if these so called ssc models fit much better the gamma ray emission in a wide energy range, there are still problems that are hard to solve. the main problem (for me) is that it is difficult to imagine a mechanism that only injects electrons and not charged nuclei in the acceleration site. one possible answer is that both electrons and cosmic rays are accelerated but the matter density around the source is very low and the cosmic rays do not interact to produce neutral mesons and gamma rays. electrons, on the other hand, emit synchrotron radiation in their propagation in magnetic fields and these low energy photons are the target for inverse compton interactions. there are, of course, several sources that fit better the hadronic interaction models and this is true for more of the combinations of supernova remnants and molecular clouds. the proof of π0 origin of the gamma rays in a source is the decrease of the flux at energies lower than 70 mev, one half of the π0 mass. this however does not happen if there is a significant contribution to the gamma ray flux from bremsstrahlung. this is one of the reasons that for most of the gamma ray sources the hadronic origin is suspected, but not proven. the suspicion is usually because of the existence of large matter density around or in front of the source. figure 5: the neutrino flux that corresponds to the gamma ray flux of ic 443. no neutrinos attributed to this sources have been observed yet. the only way we can be certain that the gamma ray generating process is hadronic is if we observe also neutrinos from the same object. gamma ray and neutrino productions are closely related as shown in eq. 1 that gives the shape of the gamma ray and neutrino fluxes as a function of the same astrophysical parameters. dnγ(ν) deγ(ν) = dncr decr fa σinel mn [ρr]cν 2znπ0(π±) γ + 1 (1) the term fa accounts for the differences between proton-proton and nuclei interactions and the term cν, that is less than 1, accounts for the different kinematics 192 are supernovae responsible for the gamma ray spectrum from the galactic center? of gamma ray and neutrino production. the gamma ray flux is higher than that of neutrinos. the big problem is the tiny neutrino interaction cross section that requires huge detectors similar to the 1 km3 icecube detector at the south pole. in some cases the low neutrino cross section is an advantage as all neutrinos generated by a source will not be absorbed and will be visible by the neutrino telescopes. acknowledgement this work is supported in part by the us department of energy contract ud-fg02-91er40626. references [1] abdo, a.a. et al. (fermi/lat collaboration): 2010, ap.j. suppl. 188, 429. doi:10.1088/0004-637x/715/1/429 [2] aharonian, f. et al. (hess collaboration): 2006, nature, 439, 695. doi:10.1038/nature04467 [3] ackermann, m (fermi/lat collaboration(: 2012, phys. rev. d85:109901; arxiv:1202.2856 [4] berezinsky, v.s., gaisser, t.k., halzen, f., & stanev, t.:1993, astropart. phys. 1, 281. doi:10.1016/0927-6505(93)90014-5 [5] bloemen, h.:1989, ann. rev. astron. astrophys. 27, 469. doi:10.1146/annurev.aa.27.090189.002345 [6] boria tridon, d. et al. (magic collaboration):2010, nim a623, 437. [7] caprioli, d., amato, e. & blasi:2010, p.:astropart. phys., 33, 160. [8] gillesen, s. et al.:2013, arxiv:1306.1374 [9] ginzburg, v.l. & syrovatskii, s.i.:1964, the origin of cosmic rays (pergamon press, oxford). [10] green, d.a.:2009, bull. astr. soc. india, 37, 45. astro-ph:0905.3699v2 [11] hanna, d. et al. (veritas collaboration):2008, nim a588, 26. [12] hinton, j.a. et al. (hess collaboration):2004, new astr. rev. 48, 331. doi:10.1016/j.newar.2003.12.004 [13] reynolds, s.p.:1998, apj 493, 375. [14] su m., slatyer t. r.& finkbeiner d. p.:2010, apj, 724, 1044. doi:10.1088/0004-637x/724/2/1044 [15] torres, d., rodriguez marrero, a.y. & de cea del pozo, e.:2010, mnras, 408, 1257. doi:10.1111/j.1365-2966.2010.17205.x discussion carlotta pittori: is there any possible correlation between your viewgraph of the galactic center with the idea of an episode of cosmic ray acceleration 106 years ago and the 511 kev integral map shown by ubertini? todor stanev: as piero ubertini said himself it is a matter of electron and positron density in the annihilation site. there is, of course, also the problem of their diffusion. how far these particles would go away from the acceleration site in a million years and how big the annihilation area would be? the hess analysis of the galactic ridge discusses the diffusion in a location where we expect very irregular and strong magnetic fields and thus fast diffusion. 193 http://dx.doi.org/10.1088/0004-637x/715/1/429 http://dx.doi.org/10.1038/nature04467 http://dx.doi.org/10.1016/0927-6505(93)90014-5 http://dx.doi.org/10.1146/annurev.aa.27.090189.002345 http://dx.doi.org/10.1016/j.newar.2003.12.004 http://dx.doi.org/10.1088/0004-637x/724/2/1044 http://dx.doi.org/10.1111/j.1365-2966.2010.17205.x introduction supernova remnants pulsar wind nebulae shell supernova remnants snr/molecular clouds other possible ideas discussion and conclusions acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0062 acta polytechnica ctu proceedings 4:62–67, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app numerical multigroup transient analysis of slab nuclear reactor with thermal feedback filip osuský∗, branislav vrban, peter ballo, štefan čerba, jakub lüley, gabriel farkas institute of nuclear and physical engineering, faculty of electrical engineering and information technology, slovak university of technology in bratislava, bratislava, slovakia ∗ corresponding author: filip.osusky@stuba.sk abstract. the paper describes a new numerical code for multigroup transient analyses with a thermal feedback. the code is developed at institute of nuclear and physical engineering. it is necessary to carefully investigate transient states of fast neutron reactors, due to recriticality issues after accident scenarios. the code solves numerical diffusion equation for 1d problem with possible neutron source incorporation. crank-nicholson numerical method is used for the transient states. the investigated cases are describing behavior of pwr fuel assembly inside the spent fuel pool and with the incorporated neutron source for better illustration of the thermal feedback. keywords: diffusion equation, neutronics, transient state, crank-nicholson method, thermal feedback. 1. introduction nuclear reactors are most reliable source of energy at present time. the advantage of new fast neutron reactor concepts is the possibility of the transmutation of the 238u into the fissionable 239pu, so the worldwide resources of fissionable materials could be enlarged when the fast neutron reactors will be commercialized. however, it is necessary to carefully investigate the transient states of these new concepts, because the recriticality after accident scenario may be achieved [1]. for this purpose, a new numerical code has been developed within the frame of neutronic calculation group at the institute of nuclear and physical engineering (inpe). the code has been used as the basic step in understanding of transient state and fast reactor r&d and we refer to it as transos. the transos code is able to solve a time dependent neutron flux distribution of slab systems with respect to the thermal feedback in multigroup approximation. the power method with infinity norm (manhattan norm) is used for a calculation of stationary state [2]. the transient state is calculated by the crank-nicholson numerical method [3]. it is possible to define the geometry in the form of coarse mesh intervals with material parameters assigned to each of the mesh interval. point and linear neutron sources can be incorporated to the design with different fluency for each investigated energy group of neutrons. the boundary conditions have to be specified in conjunction with design parameters such as temperature of the system and power generation rate. initial state of neutron flux should be given for successful run. the developed code can be used for academic purposes at the faculty and for investigation of simple systems. this paper follows the previous work [4] and next sections describe calculation methodology of time dependent and in-dependent solution, accuracy of the methodology and analysis of two nuclear systems with different 235u enrichment. higher impact of the thermal feedback can be observed on microscopic cross section properties for low energy neutron interaction with matter, so the thermal systems are investigated to show this effect. 2. calculation methodology 2.1. stationary state (time in-dependent solution) the code solves common neutron diffusion equation with internal sources for the stationary state given by [5] dg∆φg − σa,gφg − ∑ h σg→hφg+ + ∑ h σh→gφh + χg ∑ h νσf,hφh = 0. (1) lower index g in all equations represents the neutron energy group and d stands for the associated diffusion coefficient. the first term represents the leakage of neutrons from the system, the second term represents absorptions (radiation capture), the third represents scattering from the group to another group h, the fourth represents contributions of scattering from another group and the last term describes the contribution of fission source to specified energy group (χg is the proportion of neutrons emitted by fission in group g, ν is an average yield of neutrons from the fission). the box scheme was used for the discretization of examined area and therefore (1) can be written in a numerical form (2) for two energy groups 62 http://dx.doi.org/10.14311/ap.2016.4.0062 http://ojs.cvut.cz/ojs/index.php/app vol. 4/2016 numerical analysis of slab nuclear reactor with thermal feedback of neutrons [2, 4], −d̃k−1g φ k−1 g + ( d̃k−1g +d̃ k g +σ k rghk ) φk−d̃kgφ k+1 g = = λχkghk ( 2∑ g′=1 νσkfg′φ k g′ ) + hk ( 2∑ g′=1 g′ 6=g σks,g′gφ k g′ ) , (2) where σr represents the removal macroscopic cross section (3), k represents spatial node, hk is the length of the spatial interval, λ stands for the fundamental eigenvalue, σf is the fission macroscopic cross section, σs is the scattering macroscopic cross section between energy groups of neutrons, and d̃ arbitrary coupling coefficient (3) [2]. σkrg = σ k ag + 2∑ g′=1 g′ 6=g σks,gg′ d̃ k k = 2 dkg hk dk+1g hk+1 dkg hk + d k+1 g hk+1 (3) 2.2. transient state (time dependent solution) 2.2.1. crank-nicholson numerical method the crank-nicholson method is often applied to the diffusion problems and achieves good numerical stability for the transient states. the finite difference method is used as for the stationary state. this method is able to solve heat equation and similar partial differential equations [3] m ∂2u(x,t) ∂x2 = ∂u(x,t) ∂t , (4) where m represents material parameters and u is solution of the differential equation. applying the finite difference method, the central node of the common diffusion equation can be rewritten numerically for the transient state as − m̃k−1ht 2hk uk−1,j − m̃kht 2hk uk+1,j + ( 1 + m̃k−1ht 2hk + m̃kht 2hk ) uk,j = m̃k−1ht 2hk uk−1,j+1 + m̃kht 2hk uk+1,j+1 + ( 1 − m̃k−1ht 2hk − m̃kht 2hk ) uk,j+1 + f(hk), (5) where the index k represents spatial node, index j represents step time ( u·,j stands for initial condition of the function at present time step, u·,j+1 is initial condition of the function at next time step), the hk length of the spatial interval, ht length of the time interval, f(hk) is increment of the function u (particular solution). the variables and unknowns from (5) can be rewritten in the form (6) for the neutron diffusion equation (1) and (2). − d̃k−1g ht 2hk φk−1,jg − d̃kght 2hk φk+1,jg + ( 1 + d̃k−1g ht 2hk + d̃kght 2hk ) φk,jg = d̃k−1g ht 2hk φk−1,j+1g + d̃kght 2hk φk+1,j+1g + ( 1 − d̃k−1g ht 2hk − d̃kght 2hk ) φk,j+1g + χkght ( 2∑ g′=1 νσkfg′φ k,j g′ ) − σkrghtφ k,j g + ht ( 2∑ g′=1 g′ 6=q σks,g′gφ k,j g′ ) + sght, (6) where the last term s represents the external neutron source. the heat diffusion equation k ρc ∂2t(x,t) ∂x2 + q̇ = ∂t(x,t) ∂t (7) can be similarly rewritten to [6] − k̃k−1ht 2hk tk−1,j − k̃kht 2hk tk+1,j + ( 1 + k̃k−1ht 2hk + k̃kht 2hk ) tk,j = k̃k−1ht 2hk tk−1,j+ + k̃kht 2hk tk+1,j+1 + ( 1 − k̃k−1ht 2hk − k̃kht 2hk ) tk,j+1 + ( 2∑ g′=1 σkfg′φ k,j g′ ) e ht ρkck e, (8) where k̃k = 2 kk kkρkck kk+1 kk+1ρk+1ck+1 kk kkρkck + kk+1 kk+1ρk+1ck+1 , and where k represents material conductivity, ρ is material density, c stands for material heat capacity, q̇ is power generation, t stands for temperature, e is an average released energy per fission in ev and parameter e is an electron charge. accuracy of the above mentioned method is discussed in the next subsection. 2.2.2. accuracy of the crank-nicholson method the accuracy of the method for the initial and boundary conditions (more in section 3) is presented in fig. 1. the analytical solution of the differential equation for a homogenous slab system is sinus, so the problem can be approximated by function f(x) f(x) = a sin(bx) + c. 63 filip osuský, branislav vrban, peter ballo et al. acta polytechnica ctu proceedings figure 1. accuracy of crank-nicholson method. in this case, the whole system is homogeneous subcritical nuclear system. it is necessary to insert the external neutron source because the neutron flux will never stabilize on non-zero value without it. the external neutron source is linear, located between 217–227 mm, with fluency 1.0 × 106 s−1 in thermal energy spectrum. deviations between the modeled problem and analytical solution are presented in table 1. parameter mean value dispersion rel. error a 111 935 126.1 0.1127 % b 7.076 99 × 10−3 2.015 × 10−6 0.028 47 % c −785.941 97.17 12.36 % table 1. deviation between numerical and analytical solution. amplitude (a) and shape parameter (b) are in accordance with the analytical solution (table 1). deviance in location parameter (c) is caused by the external neutron source and by zero flux boundary condition. 3. geometry of the nuclear slab system the investigated geometry represents middle cut of one heterogenic vver440 fuel assembly that is placed in the middle. one edge of the investigated geometry is represented by homogenous representation of the heterogeneous fuel assembly. control rod assembly is placed on the opposite edge. this case was introduced due to simulation of the transient state during operational conditions. geometry is also simplified into one dimension and some lengths are changed for faster convergence of numerical calculation (fig. 2). red material represents uo2 (dimension d1 is 7 mm and is the same for every red and orange material). orange material is the representation of uo2 with 3.35 % of 153gd. the two cases of the enrichment are investigated (table 2). green is b4c (dimension d5 is 142 mm), figure 2. graphical representation of investigated nuclear reactor core. enrichment material case 1 case 2 uo2 5 % 22 % uo2 + 153gd 4.4 % 21.4 % table 2. enrichment of uo2. blue stands for h2o (dimension d2 is 5 mm, d3 = 19 mm, d4 = 10 mm). purple is a mixture of the whole area between x–y points (d5 is 142 mm). all cross sections are obtained from endf/b-vii.1 library for the temperatures of 296.3 k, 600 k and 1800 k, for the discrete energies of 0.0253 ev and 2 mev [7]. isotropic scattering is assumed and the diffusion coefficient can be calculated according to (9) [8], where σt stands for the total macroscopic cross section, σtr is the transport macroscopic cross section, and µ̄0 represents unsymmetrical parameter of neutron scattering (if 64 vol. 4/2016 numerical analysis of slab nuclear reactor with thermal feedback isotropic scattering is applied µ̄0 = 0). in this case, a molecular cross section of h2o is not considered, even if the molecular cross section is higher than calculated cross sections of particular nuclides [8]. d ' λtr 1 3 = 1 3σtr = 1 3 ( σt − µ̄0σs ) ≈ 1 3σt . (9) more details about geometry input can be found in [4]. it is necessary to note that this is the only theoretical approximation and that it is mainly because the values of the macroscopic scattering cross section were simplified with certain degree of uncertainty. 4. initial and boundary conditions the zero incoming current boundary condition is used in all cases of the neutron diffusion α = j φ = 0.5, where j is the neutron current and φ is the neutron flux. this can be interpreted as that the environment behind the boundary of the nuclear system is vacuum. according to (2), the stationary state is represented by proportional neutron flux distribution that is normed to maximal value 1. it is necessary to set the initial condition in the form of total power generation p = ∫ v e ( 2∑ g′=1 σfg′φg′ ) e dv. (10) the absolute value of the neutron flux distribution is calculated from the proportional neutron flux distribution (2) and from the initial power generation condition (10). the power generation rate is calculated for the case 1 from the typical power generation of vver reactor (1375 mwth) and it is approximately 1.576 kw/m per the assembly p [w/m] = pth assembly nr · height = = 1375 × 106 349 · 2.5 = 1576, where assmebly nr is the total number of fuel assemblies in the reactor and height represents height of the core. the power generation rate is the same for the case 2 with higher enrichment to lower impact of the thermal feedback to ensure the nuclear system’s subcritical state. temperature is kept at constant 296.3 k at the boundaries that correspond to dirichlet boundary condition. temperature of the whole system is also set to 296.3 k before the transient for all cases. no thermo-mechanical physics is applied (the densities of the particular materials are constant). also no fluid mechanics is applied (the heat diffusion equation solves just thermal conduction equation) and this simplification results in total temperature increase of the whole nuclear system. 5. results 5.1. case 1 keff = 0.282 601 is calculated from the stationary state of trnasos code for this scenario. this condition may be achieved in spent fuel pool. the worst case is simulated by the neutron flux corresponding to the operational state before the transient (the power generation is 1.576 kw/m). it is possible to observe (fig. 3) that the thermal energy spectrum of neutrons decreases rapidly due to deep subcritical state of the nuclear system. the same evolution can be seen in case of fast energy spectrum of neutrons in fig. 4. the temperature is set to 296.3 k at the beginning of the transient state. the power is generated by (10) according to the neutron flux and the temperature evolution is shown in fig. 5. the temperature increases up to 460 k and according to the temperature change, the microscopic cross sections are modified. the whole system is below melting point temperature of the fuel. figure 3. neutron flux distribution for energy group 0.0253 ev in case 1 (the neutron flux decreases with time evolution due to deep subcritical state). figure 4. neutron flux distribution for energy group 2 mev in case 1 (the neutron flux decreases with time evolution due to deep subcritical state). 65 filip osuský, branislav vrban, peter ballo et al. acta polytechnica ctu proceedings figure 5. temperature evolution in case 1 with keff = 0.282601 (the temperature decreases with time evolution – the nuclear system is brought into safeconditions). 5.2. case 2 the keff is equal to 0.982 693 in this case, due to higher enrichment of 235u (table 2). the neutron source is incorporated within the core and its total emission of neutrons is 1.1 × 106 s−1 in the area between 217–227 mm on x-axis. it corresponds to the concept of super-thermal liquid-helium (4he) source (ucn) [9] and portion of 1.0 × 106 s−1 is in the thermal energy spectrum. subcritical multiplication can be derived from n∞ = s 1 1 −keff = 1.1 × 106 1 1 − 0.982 693 ≈ 108. the result of subcritical multiplication formula is consistent with the calculated results that are shown in fig. 6 and fig. 7. the whole system starts from zero power level. the transient state of the neutron flux is stabilized approximately after 50 000 s. however, the transient of the temperature evolution is represented by higher inertia and the transient state is stabilized in the region of 140 000 s (fig. 8). the temperature does not exceed 380 k. figure 6. neutron flux distribution for energy group 0.0253 ev in case 2 (nuclear system is stabilized after subcritical multiplication for keff = 0.982 693). figure 7. neutron flux distribution for energy group 2 mev in case 2 (nuclear system is stabilized after subcritical multiplication for keff = 0.982 693). figure 8. temperature evolution in case 2 (temperature is stabilized below melting point of fuel material for keff = 0.982 693). 6. conclusions the successful application of the transos code was demonstrated. it has been found that the maximum value of the neutron flux and its temperature is located in the middle of the system within the heterogeneous region. on the other hand, the diffusion theory is more appropriate for the homogenous problems and calculation error increases with greater heterogeneity. in all cases, the temperature does not exceed melting point of fuel material. the code that was written in c++ can be used for academic purposes at the institute and for solution of simple nuclear systems. it has been confirmed that the diffusion theory is suitable for fast calculations and it is used during development of theoretical reactor design. therefore, development of simple codes increases the knowledge level of neutron physics and it can result in enhancing of nuclear safety. the contribution of 238u thermal feedback with thermo-hydraulic properties of the system has to be investigated in the future. acknowledgements this work was financially supported by grant of science and technology assistance agency no. apvv-0123-12 and stu grant scheme for support of young researchers. 66 vol. 4/2016 numerical analysis of slab nuclear reactor with thermal feedback references [1] h. ninokata, t. sawada, h. tomozoe, et al. a study on recriticality characteristics of fast reactors in pursuit of recriticality-accident-free concepts. progress in nuclear energy 29:387–393, 1995. doi:10.1016/0149-1970(95)00067-t. [2] j. han-gyu. solution of one-dimensional, one-group neutron diffusion equation, lecture note 1. seoul national university, reactor physics laboratory, 2008. [3] p. b. patil, u. p. verma. numerical computational methods. alpha science international, ltd, 2006. [4] f. osuský, b. vrban, š. čerba, et al. two-group numerical analysis of one dimensional table nuclear reactor. in proceedings of the 20th international conference on applied physics of condensed matter (apcom2014), pp. 320–323. 2014. [5] p. reuss. neutron physics. edp sciences, 2008. [6] m. kalousek, b. hučko. prenos tepla. vydavateľstvo stu, bratislava, 1996. [7] nuclear data services. point2015 data. temperature dependent endf/b-vii.1 cross section data, https://www-nds.iaea.org/point/. [8] j. r. lamarsh. introduction to nuclear reactor theory. amer nuclear society, 2002. [9] e. lychagin, v. mityukhlyaev, a. muzychka, et al. ucn sources at external beams of thermal neutrons. an example of pik reactor. nuclear instruments and methods in physics research section a: accelerators, spectrometers, detectors and associated equipment 823:47–55, 2016. doi:10.1016/j.nima.2016.04.008. 67 http://dx.doi.org/10.1016/0149-1970(95)00067-t https://www-nds.iaea.org/point/ http://dx.doi.org/10.1016/j.nima.2016.04.008 acta polytechnica ctu proceedings 4:62–67, 2016 1 introduction 2 calculation methodology 2.1 stationary state (time in-dependent solution) 2.2 transient state (time dependent solution) 2.2.1 crank-nicholson numerical method 2.2.2 accuracy of the crank-nicholson method 3 geometry of the nuclear slab system 4 initial and boundary conditions 5 results 5.1 case 1 5.2 case 2 6 conclusions acknowledgements references 96 acta polytechnica ctu proceedings 1(1): 96–102, 2014 96 doi: 10.14311/app.2014.01.0096 star formation in tadpole galaxies casiana muñoz-tuñón1,2, jorge sánchez almeida1,2, debra m. elmegreen3, bruce g. elmegreen4 1instituto de astrof́ısica de canarias, e-38205 la laguna, tenerife, spain 2departamento de astrof́ısica, universidad de la laguna, tenerife, spain 3department of physics and astronomy, vassar college, poughkeepsie, ny 12604, usa 4ibm research division, t.j. watson research center, yorktown heights, ny 10598, usa corresponding author: cmt@iac.es abstract tadpole galaxies look like a star forming head with a tail structure to the side. they are also named cometaries. in a series of recent works we have discovered a number of issues that lead us to consider them extremely interesting targets. first, from images, they are disks with a lopsided starburst. this result is firmly established with long slit spectroscopy in a nearby representative sample. they rotate with the head following the rotation pattern but displaced from the rotation center. moreover, in a search for extremely metal poor (xmp) galaxies, we identified tadpoles as the dominant shapes in the sample – nearly 80% of the local xmp galaxies have a tadpole morphology. in addition, the spatially resolved analysis of the metallicity shows the remarkable result that there is a metallicity drop right at the position of the head. this is contrary to what intuition would say and difficult to explain if star formation has happened from gas processed in the disk. the result could however be understood if the star formation is driven by pristine gas falling into the galaxy disk. if confirmed, we could be unveiling, for the first time, cool flows in action in our nearby world. the tadpole class is relatively frequent at high redshift – 10% of resolvable galaxies in the hubble udf but less than 1% in the local universe. they are systems that could track cool flows and test models of galaxy formation. keywords: starburst galaxies tadpoles cool flows rotation curves galaxy disks. 1 introduction elongated galaxies with bright clumps at one end are visible in deep field images taken with hst or from the ground; van der bergh et al., 1996 called them “tadpole” galaxies”. figure 1 in elmegreen et al. (2005) shows different morphologies of galaxies observed with the hubble ultra deep field (udf). the fourth row presents images of tadpoles. this asymmetric morphology is rather common at high redshift but rare in the local universe. for example, tadpoles constitute 10% of all galaxies larger than 10 pixels in the udf (elmegreen et al. 2007; elmegreen & elmegreen, 2010), and they represent 6% of the udf galaxies identified by straughn et al. (2006) and windshorst et al. (2006) using automated search algorithms. in contrast, elmegreen et al. (2012; hereafter paper i), find only 0.2% tadpoles among the uv-bright local galaxies of the kiso survey by miyauchi-isobe et al. (2010). this decrease suggests that the tadpole morphology represents a common but transition phase during the assembly of some galaxies. since local tadpole galaxies are very low mass objects compared to their high redshift analogues, this phase must be already over for the local descendants of high redshift tadpoles. the tadpole structure has inspired several explanations, such as ram pressure stripping that triggers star formation at the leading edge, to mergers (see sánchez almeida et al. 2013 for an extensive review). the explanation that we propose, based on observational evidence summarized here, is that the starburst head may result from the accretion of an external flow of pristine gas that penetrates the dark matter halo and hits and heats a pre-existing disk, which is viewed to the side as the tail. we will briefly present a summary of recent results showing, first that the local tadpoles share the properties of their higher redshift and higher mass counterparts (section 2), that they belong to the extremely metal poor (xmp) sample of the blue compact dwarfs family (section 3), they are rotating discs (section 4) and, finally, that the starbursts (heads) show a drop in the already low metallicity that can only be understood if fresh metal-poor gas is falling onto the galaxy (section 96 http://dx.doi.org/10.14311/app.2014.01.0096 star formation in tadpole galaxies 5). we finish with a brief summary and future actions. 2 local and hight z tpg share properties we used sloan digital sky survey data to determine the ages, masses, and surface densities of the heads and tails in 14 local tadpoles (shown in figure 1) selected from the kiso and michigan surveys of uv-bright galaxies, and we compared them to tadpoles previously studied in the hubble ultra deep field. the result, published in paper i is that the young stellar mass in the head scales linearly with the rest-frame galaxy luminosity, ranging from ≈ 105 m� at galaxy absolute magnitude u = −13 mag to 109 m� at u = −20 mag. the tails in the local sample look like bulge-free galaxy disks. their photometric ages decrease from several gyr to several hundred myr with increasing redshift. the far-outer intensity profiles in the local sample are symmetric and exponential. we suggest that most local tadpoles are bulge-free galaxy disks with lopsided star formation, perhaps from environmental effects such as ram pressure or disk impacts, or from in-situ gas collapse to a giant star-forming clump with a jeans-length that is comparable to half the disk size. the existence of a disk, proposed from the analysis of the luminosity profiles, has been further confirmed spectroscopically (see section 4). figure 1: tadpole galaxies from the kiso and um samples. figure from elmegreen et al., 2012. 3 belong to the bcdsxmp class blue compact galaxies are important targets for a number of reasons. they are small systems, thought to remain as fossils or debris from the formation of larger galaxies in the early epochs of the universe. their low metallicities, high ssfrs and low doubling time (less than 1gy) and active stabursts make them ideal targets to study the old universe at low redshift. their complete census, as well as the likelihood that they have a long quiescent phase between starbursts, took us to study the whole sample by making use of sdss. the properties of the quiescent and burst phases were determined from detailed studies of the bursts and host galaxies of a nearby sample (amorin et. al., 2007, 2008). these properties allowed us to identify such objects in the database. the sdss dr6 database provides ≈ 21.500 quiescent bcd candidates, a number 30 times larger to those bursting (bcds). this result implies that one out of every three dwarf galaxies in the local universe may be a quiescent bcd. the properties of the two samples are consistent with a single sequence in galactic evolution with the quiescent phase lasting 30 times longer than the burst phase (sánchez almeida et. al., 2008). in sánchez almeida et al. (their figure 9) there is a clear subsample of objects with the lowest luminosity which turn out to be also those that are most metal poor. we carried out a systematic search for extremely metal-poor (xmp) galaxies in the spectroscopic sample of sloan digital sky survey (sdss) data release 7 (dr7) (morales-luis et al., 2011). the xmp candidates are found by classifying all of the galaxies according to the form of their spectra in a region 80 å wide around hα using an automatic classification algorithm, k-means (sánchez almeida et al.,2009). our systematic search renders 32 galaxies having negligible [n ii] lines, as expected in xmp galaxy spectra. twenty-one of them were previously identified as xmp galaxies in the literature, and 11 were new. this was established after a thorough bibliographic search that yielded only some 130 galaxies known to have an oxygen metallicity 10 times smaller than the sun (explicitly, with 12 + log [o/h] ≤ 7.65). xmp galaxies are rare; they represent 0.01% of the emission lines galaxies in sdss/dr7. the xmp galaxies constitute 0.1% of the galaxies in the local volume, or ≈ 0.2% of the emission-line galaxies. all but four of our candidates are bcd galaxies, and 24 of them have either cometary shape or are formed by chained knots. note that this result, the morphology, is absolutely independent of our search criterium (spectra). 4 tadpoles are rotating disks the work in paper i was followed up by mean of high spectral resolution long slit observations at the int (2.5m) at observatorio del roque de los muchachos (orm) with ids spectrograph. in order to determine the dynamical properties and metallicities of lo97 casiana muñoz-tuñón et al. figure 2: sdss mugshots of all xmp candidates (from spectroscopy). the figure is taken from morales-luis et al., 2011. note the prevalence of tadpole shapes. cal tadpoles, we measured hα spectra along the headtail direction in a representative fraction of the original sample. further details are in sánchez almeida et al. (2013). in figure 3, we show spectral flux (the solid lines) and hα flux (the dashed lines) along the slit. note the obvious lopsidedness of the light distributions, as expected from the tadpole shape. the origin of distance on the abscissa has been set as the position of the maximum of the spectral flux distribution. the dotted line represents a gaussian fitted to the hα flux around the head. the horizontal bar in each panel gives a common length scale corresponding to 1 kpc. the thin vertical solid lines indicate the center of rotation obtained from the rotation curve fit shown in fig. 4. velocities, masses, abundances and other physical parameters were determined from the spectra. bulk velocities were measured from the displacement of hα. we computed the displacement both as the barycenter of the emission line, and as the center of a gaussian function fitted to the profile. errors were estimated from the s/n measured in the continuum and then propagated to the centroids. the fwhm of the profiles were also measured directly from the profiles and from the gaussian fits. their errors were also inferred from the noise measured in the continuum by error propagation. figure 4 shows the velocity curves of the tadpole galaxies included in the study. the abscissae represent distances along the major axes of the galaxies from the position of the tadpole head; i.e., the brightest point on the galaxy. the range of distances differs for the different targets, but the horizontal bar in each panel gives a common length scale corresponding to 1 kpc. the points with error bars show the observations whereas the thick solid line represents the best fit of the observed points to the analytic rotation curve. the part of the rotation curve shown in red indicates the portion of the velocity curve used for fitting. the thin horizon98 star formation in tadpole galaxies tal and vertical lines indicate the systemic velocity and the center of rotation obtained from the fit, respectively. the little arrows on top of each panel also indicate the centers of rotation. the zero of the velocity scale is set by velocity of the spatially integrated spectrum of the galaxy; positive velocities are redshifts. figure 3: spectral flux (the solid lines) and hα flux (the dashed lines). formal error bars for photometry are not included since they are negligibly small. the main result shown in figure 4 is that five out of seven targets show velocity gradients interpreted as rotation. another important result is that the tadpole head is not at the rotation center. figure 4: velocity curves of the seven tadpole galaxies obtained from long slit spectroscopy. figure from sánchez almeida et al. 2013 sometimes the interpretation of the velocity curve as rotation is obvious (e.g., kiso8466), but other times the curve looks more like a perturbed rotation (e.g., kiso5639). kiso3193 and kiso3867 have a rather flat velocity curve, and therefore no obvious rotation. however, one of them, kiso3867, shows a systematic line shift of the order 10–20 km s−1 between the two extremes of the galaxy (figure 4). the amplitude is of the order of the error bars, but the displacement is in the raw data as judged by inspection of the individual hα profiles. 5 drops in metallicity at the head location the metallicity along the slit was estimated using the ratio [nii]λ6583 to hα. figure 5 taken from sánchez almeida et al (2013) shows the variation across the galaxies of the oxygen abundance, including their error bars. all galaxies have sub-solar metallicity the thick horizontal line marks the solar oxygen abundance given by 12 + log(o/h)� = 8.69 ± 0.05 (asplund et al., 2009). the galaxies also present significant abundance gradients, with the lowest abundances tending to coincide with the largest hα emissions (e.g., kiso6669 and kiso6877) in fig. 5. remember that the vertical dotted lines mark the position of the peak hα fluxes. two targets, kiso5639 and kiso6877, have metallicities well below one-tenth the solar value, therefore, they belong to the xmp galaxies class (kunth & östlin, 2000). xmp galaxies are really rare objects: one out of a thousand galaxies in the local universe according to morales-luis et al. (2011). therefore the fact that we observe two in a sample of seven cannot be a coincidence. it is known that a significant fraction of xmp galaxies turn out to be cometary or tadpole (papaderos et al., 2008). we have found that the reverse holds too, i.e., that tadpole galaxies have a significant probability of being xmp. this fact supports the idea that the tadpole morphology is a sign of dynamical youth, as the low metallicity is a sign of being chemically young (see discussion in sánchez almeida et al., 2013). 6 summary and future tadpole galaxies have an easily identified shape, belong to the blue compact dwarf class, and are rare in the nearby universe. their appearance suggests then to be the consequence of some interaction or star formation triggered by ram pressure processes. these local tadpoles seem though to form a continuous sequence with the udf tadpoles, seen in relatively larger number at high redshift (elmegreen & elmegreen 2010). regarding their photometric properties, local tadpoles occupy the low mass end in sequences such as star formation, surface density and mass-to-light ratio. in addition, the radial intensity profiles of both samples (local and high redshift tadpoles) show an exponential decrease at large galactocentric distances, which we interpreted as an evidence for the existence of an underlying disk. from the point of view of their chemical content, extremely metal poor (xmp) galaxies are the least evolved objects in the local universe (pagel et al., 1992; 99 casiana muñoz-tuñón et al. figure 5: oxygen abundance variation across the galaxies. the vertical solid line represents the center of rotation, whereas the vertical dotted line indicates the location of the maximum in the hα flux profile. the thick horizontal solid line indicates the solar metallicity. note the existence of abundance variations, with the minima coinciding with the largest hα signals. note also that kiso5639 and kiso6877 reach very low abundances, below one-tenth of the solar value; therefore, they are members of the set of rare xmp galaxies. kunth & östlin, 2000). they represent only 0.1 % of the galaxies in an arbitrary nearby volume (moralesluis et al., 2011) and a significant fraction of these chemically primitive objects turn out to have tadpole or cometary shape (brinchmann et al. 2008). this association between low metallicity and tadpole shape suggests that both are attributes of very young systems. detailed studies of a significative sample of the local tadpole class by means of high-resolution long-slit spectroscopy put into evidence two important facts: they do rotate and they show oxygen abundances that vary along the disk. a high percentage of them are xmp galaxies and the metallicity is minimum were starformation is maximum. moreover, the rotation center does not coincide with the current starburst location (the head). the oxygen metallicity estimated from [nii]6583/hα often shows significant spatial gradients across the galaxies (∼0.5 dex), being lowest at the head and increasing in the rest of the galaxy, tail included. a similar result with the highest sfr been the lowest metallicity region has also been found in low metallicity grb host galaxies (levesque et al., 2011). the sense of the resulting metallicity gradient differs from the observation of local disk galaxies, where the gas-phase metallicity increases toward the galaxy centers (vilchez et al., 1988) or is just constant (genel et al., 2008). however, the type of variation we measure, with a minimum metallicity at the most intense star-forming region, has been observed in galaxies at redshift around 3 by cresci et al. (2010), where it is interpreted as evidence for infall of pristine gas triggering star formation. as a result of all the evidence found, we propose that local tadpole galaxies are disks still in the process of being formed in our present-day universe. moreover 100 star formation in tadpole galaxies their otherwise very low star formation is enhanced and triggered by cool-flow accretion of pristine metal-poor gas. this triggering causes an extremely metal-poor head to shows up on the side of an immature (still forming) rotating disk. there are a number of parallel actions now being undertaken to go further. we have just finished a study of the hi content of a complete sample of xmp galaxies (fihlo et al., 2013) and compared the properties of the clumps of different galaxy types at different redshifts (elmegreen et al., 2013). the recent analysis of a new sample corroborates the metallicity drop in the xmp galaxies using the so call direct method (sánchez almeida et al. 2014). cosmological simulations predict cold-flow buildup to be the main mode of galaxy formation (dekel et al., 2009). the incoming gas is expected to form giant clumps that spiral in and merge into a central spheroid. to confirm such a scenario at high redshift is neither easy nor feasible and we think we have discovered nearby cases that may allow for comparisons. we are making a comprehensive study of the disks and bursts of other bcds with xmp properties to search for extremely metal poor clumps in 2d spectroscopy. the final aim is to characterize the local xmp galaxy sample within the cool flow paradigm. acknowledgement this work has been partly funded by the spanish ministry for science, project aya 2010-21887-c04-04. results based on observations at the observatorio del roque de los muchachos (orm), operated by the iac at la palma. thanks to franco giovannelli and lola sabau-graziatti, organizers of this workshop for their kindness and dedication. thanks to the referee for useful comments. references [1] van den bergh, s., abraham, r.g., ellis, r.s., tanvir, n.r., santiago, b.x., & glazebrook, k.g. 1996, aj 112, 359 [2] emegreen, d.m., elmegreen, b.g., rubin, d.s. & schaffer, m.a. 2005, apj, 631, 85 doi:10.1086/432502 [3] elmegreen, d. m., elmegreen, b. g., ravindranath, s., & coe, d. a. 2007: apj, 658, 763 doi:10.1086/511667 [4] elmegreen, b. g. & elmegreen, d. m. 2010: apj, 722, 1895 doi:10.1088/0004-637x/722/2/1895 [5] straughn, a. n., cohen, s. h., ryan, r. e., et al. 2006, apj, 639, 724 doi:10.1086/499576 [6] windhorst, r. a., cohen, s. h., straughn, a. n., et al. 2000, newar, 50, 821. 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[26] sánchez almeida, j., morales-luis, a.b, muñoztuñón, c., et al., 2014: apj in press. 102 http://dx.doi.org/10.1038/nature07648 introduction local and hight z tpg share properties belong to the bcdsxmp class tadpoles are rotating disks drops in metallicity at the head location summary and future 157 acta polytechnica ctu proceedings 1(1): 157–162, 2014 157 doi: 10.14311/app.2014.01.0157 agile highlights after six years in orbit carlotta pittori1,2 on behalf of the agile collaboration 1asi science data center (asdc), via del politecnico snc, 00133 roma, italy. 2inaf-oar, via frascati 33, i00040 monte porzio catone (rm), italy. corresponding author: carlotta.pittori@asdc.asi.it abstract agile is an asi space mission in collaboration with inaf, infn and cifs, dedicated to the observation of the gammaray universe in the 30 mev – 50 gev energy range, with simultaneous x-ray imaging capability in the 18–60 kev band. the agile satellite was launched on april 23rd, 2007, and produced several important scientific results, among which the unexpected discovery of strong flares from the crab nebula. this discovery won to the agile pi and the agile team the bruno rossi prize for 2012 by the high energy astrophysics division of the american astronomical society. thanks to its sky monitoring capability and fast ground segment alert system, agile detected many galactic and extragalactic sources: among other results agile discovered gamma-ray emission from the microquasar cygnus x-3, detected many bright blazars, discovered several new gamma-ray pulsars, and discovered emission up to 100 mev from terrestrial gamma-ray flashes. we present an overview of the main agile data center activities and the agile scientific highlights after 6 years of operations. keywords: gamma-rays: observations catalogs. 1 introduction agile (astro-rivelatore gamma a immagini leggero) (tavani et al., 2009a) is a gamma-ray astrophysics mission of the italian space agency (asi), with inaf, infn and cifs participation which has been in orbit since 23 april 2007. the agile payload is composed of a pairproduction gamma ray imager (grid) sensitive in the energy range 30 mev-50 gev (barbiellini et al., 2001; prest et al., 2003), an x-ray imager (superagile) sensitive in the energy range 18-60 kev (feroci et al., 2007), and a mini-calorimeter (mcal) sensitive to gamma-rays and charged particles with energies between 300 kev and 100 mev (labanti et al., 2009). the instrument, weighting only ∼ 100 kg, is the most compact ever operational for high-energy astrophysics (approximately a cube of about 60 cm size) with excellent detection and imaging capability. agile is characterized by a very large field of view (∼ 3 sr), a good angular resolution, 0.1 0.2 degrees in gamma rays and 1-2 arcminutes in x rays, and a small dead time (100 µs), which make it a very good instrument to study persistent and transient gamma-ray sources even on very short timescales. recent calibration results of the agile-grid using in-flight data and monte carlo simulations producing instrument response functions have been presented in chen et al. (2013). 2 the ground segment and the agile data center the agile data center (adc) is part of the asi science data center (asdc), previously located in the esa establishment of esrin in frascati, and recently moved within the asi headquarters, rome (italy). satellite data are routinely collected every ∼ 96 minutes by the asi ground station in malindi (kenya), then quickly sent to the satellite operations centre in telespazio, fucino, and then transferred, preprocessed, analyzed, and stored at the adc. adc is in charge of all the scientific oriented activities related to the analysis, archiving and distribution of agile data. from july 2007 july to october 2009 (∼ 2.5 years) agile was operated in “pointing observing mode”, characterized by long observations called observation blocks (obs), typically of 2-4 weeks duration, following a predefined baseline pointing plan, mostly concentrated along the galactic plane. on november 4, 2009, agile scientific operations were reconfigured following a malfunction of the rotation wheel occurred in mid october, 2009. the satellite is currently operating regularly in “spinning observing mode”, surveying a large fraction (about 70%) of the sky each day. since december 21, 2012 due to asi’s malindi ground station technical problems, the acquisition of telemetry data from the agile satellite has been reduced. 157 http://dx.doi.org/10.14311/app.2014.01.0157 carlotta pittori figure 1: whole sky agile intensity map (ph cm−2 s−1 sr−1) in galactic coordinates and aitoff projection, for energies e > 100 mev, accumulated during the first ∼ 6 years of observations up to december, 2012 (pointing + spinning observing modes). malindi downlink efficiency for agile-grid data is currently at a daily level of about 80%. all agile payload functions are nominal, and the daily grid skymonitoring activities are fully operational, despite the telemetry gaps in the archive due to intermittent data acquisition at the ground station. the agile quicklook (ql) alert system is composed of two independent automated analysis parts: the ql scientific pipeline running at adc (pittori et al., 2013) and the agile-grid science alert system (sas) pipeline running at the inaf-iasf bologna (bulgarelli et al., 2013). data are automatically analyzed at every downlink and accumulated on different timescales. lightcurves of potential gamma-ray sources are produced using blind search techniques and crosscorrelation with a reference list of known gamma-ray emitters. as proper flux thresholds are exceeded, alerts for transient gamma-ray sources are automatically generated and notified through e-mails, sms messages, and also through a dedicated application for smartphones and tablets. these alerts are crosschecked, and a manual analysis is performed for the most interesting candidate gamma-ray sources. the daily monitoring activity of the agile ql alerts resulted in the publication of 104 atel and 40 gcn up to may, 2013, among which we mention the first detection of gamma-ray emission from microquasars above 100 mev, from cygnus x-3 and cygnus x-1 in the cygnus region, the prompt alert of many gamma-ray flares from blazars and galactic transients, and the surprising discovery of gamma-ray flares from the crab nebula. the 2012 bruno rossi international prize has been awarded to the pi, marco tavani, and the agile team for this important and unexpected discovery. 3 new agile catalogs and asdc interactive web tools the agile first catalog of high-confidence γ-ray sources detected by agile during its first year of operations (pittori et al., 2009) was recently updated using data covering the whole period of pointed observations, ∼ 2.3 years, in verrecchia et al. (2103) (“an updated list of agile bright gamma-ray sources and their variability in pointing mode”). refined analysis of complex regions of the galactic plane with improved event filter and calibrations yielded a new reference list, and a multi-source maximum likelihood analysis was performed over the timescale of the agile pointed observation blocks (ob) resulting in an updated list of 54 sources. eight 1agl sources, with significance above 4σ from the analysis over the entire observing period, were not detectable on the ob timescale either due to low ob exposures and/or to their positions within complex galactic regions. the new agile-grid variability catalog (verrecchia et al., 2103) is also available from the adc webpages as an interactive web table1. in the on-line version, light curves from fluxes in each ob are accessible within the asdc data explorer, 1http://www.asdc.asi.it/agile1rcat/ 158 agile highlights after six years in orbit in the agile-grid data products tab. the asdc data explorer is a multimission web tool which allows the user to browse both internal multimission archive catalogs (grouped by energy band) and external catalogs from other services (vizier, ned, simbad etc.), providing an easy way to explore and cross-correlate large data sets from radio to tev. moreover by using the asdc sed builder tool (v3.0) it is possible to build and handle spectral energy distributions (seds), time resolved seds, and multi-frequency lightcurves. virtual observatory (vo) tools like topcat can also be used to handle sed data in the time domain. the agile reference list of bright gamma-ray sources above 100 mev over the first 2.3 years of observations thus includes at the moment 62 sources, as compared to the 47 1agl in the first catalog. work is in progress to build new complete catalogs of agilegrid sources over the whole observing period (pointing+spinning). the mcal instrument of the agile satellite can observe the high-energy part of gamma-ray bursts (grb) from 350 kev to 100 mev with good timing capability, and the agile-mcal grb catalog, including 85 grb observed from april 2007 to october 2009, was recently published in galli et al. (2013). an online version of the new agile-mcal grb catalog is available from the adc web pages2. we adapted the asdc data explorer to provide the user with additional services dedicated to gamma-ray bursts through a grb explorer tool. the interactive web table in this case contains all the data for the published grbs, including spectral data for 21 bursts, plus some supplementary information on other 7 events. light curves and energy spectra (when available) can be accessed from the grb explorer tool, in the agile mcal data products tab. it is also possible to browse external services to look for grb-related data products from other missions, such as swift-xrt light curve from leicester repository, swift-bat product analysis, fermigbm quicklook lightcurve, gcns and blog for the chosen grb. 4 agile science highlights fig. 1 shows the total gamma-ray intensity above 100 mev as observed by agile up to december, 2012, during the first ∼ 6 years of observations. we present here a selection of the main agile science highlights and some recent updates. agile has provided the first detection of a colliding wind binary (cwb) system above 100 mev in the η-carinae region (tavani et al., 2009b). agile detected a gamma-ray source (1agl j1043-5931, now 1aglr j1044-5944) consistent with the position of the cwb massive system η-car during the time period 2007 july 2009 january. a 2-day gamma-ray flaring episode was also reported on 2008 oct. 11-13, possibly related to a transient acceleration and radiation episode of the strongly variable shock in the system. a revised gamma-ray source list in the complex carina region shown in fig. 2 has been published in verrecchia et al. (2013). figure 2: agile intensity map of the carina region above 100 mev over ∼ 2.3-year pointing mode observation period. new sources and revised (r) positions for 1agl sources are indicated in magenta, the 1agl error circles are indicated in black. the agile detection of gamma-ray emission from the pwn vela-x, described in a science paper (pellizzoni et al., 2010), has been the first experimental confirmation of gamma-ray emission (e> 100 mev) from a pulsar wind nebula. this result constrains the particle population responsible for the gev emission and establishes a class of gamma-ray emitters that could account for a fraction of the unidentified galactic gamma-ray sources. subsequently the nasa fermi satellite has confirmed the vela-x gamma-ray detection, and has also firmly identified other 4 pulsar wind nebulae plus a large number of candidates. gamma-ray flaring activity for a source positionally consistent with cygnus x-1 microquasar was reported twice by agile (bulgarelli et al., 2010; sabatini et al., 2010a and 2010b). agile extensive monitoring of cygnus x-1 in the energy range 100 mev 3 gev during the period 2007 july 2009 october confirmed the existence of a spectral cutoff between 1-100 mev during the typical hard x-ray spectral state of the source. however, even in this state, cygnus x-1 is capable of producing episodes of extreme particle acceleration on 1-day timescales, and even shorter lived flares in the tev range as detected by magic in 2006 (albert et al., 2007). gamma-ray flares of cygnus x-1 detected 2http://www.asdc.asi.it/mcalgrbcat/ 159 carlotta pittori by agile have been recently confirmed for the first time by a reanalisis of fermi-lat data (bodaghee et al., 2012). agile detected for the first time several gammaray flares from cygnus x-3 microquasar and also a weak persistent emission above 100 mev from the source (tavani et al., 2009c). gamma-ray flares occur either in coincidence with low hard-x-ray fluxes or during transitions from low to high hard-x-ray fluxes, and usually appear before major radio flares, following a clear repetitive pattern. this important agile finding has been confirmed by fermi-lat, which also detected the orbital period (4.8 hours) of the binary system in gammarays, securing unambigously the temporal signature of the microquasar (abdo et al., 2009). in the 9 days from december 2 to december 11, 2009 agile and fermi were able to answer a long-lasting question: cygnus x3 system is able to accelerate particles up to relativistic energies and to emit gamma-rays above 100 mev. figure 3: the cygnus region in gamma-rays: agile intensity map from 100 mev to 10 gev. data taken in ∼ 2-year pointing observing mode, from nov. 2007 to oct. 2009. figure adapted from g. piano presentation, 9th agile workshop, 2012. understanding the origin of cosmic rays is one of the most important issues of high-energy astrophysics, and galactic supernova remnants (snr) are considered to be an ideal laboratory to study cosmic-ray acceleration. experimental data analysis below e = 200 mev is crucial to discriminate between theoretical models, since in this energy range hadronic and leptonic emission spectra have a well distinct behavior due to a steepening of the hadronic spectrum due to the neutral pion emission, which is missing in the leptonic case. the agile gamma-ray imager reaches its optimal sensitivity just at the energies in the 50 mev a few gev range at which neutral pions produced by proton-proton interactions radiate, and it was the first to discover a pattern of gamma-ray emission from the supernova remnant w44 that, combined with the observed multifrequency properties of the source, can be unambiguously attributed to accelerated protons interacting with nearby dense gas. this important agile result was reported in (giuliani et al., 2011), and recently confirmed by new fermi-lat data (ackermann et al., 2013) before agile and fermi observations a direct identification of sites in our galaxy where proton acceleration takes place was elusive. figure 4: the w44 supernova remnant spectral energy distribution in gamma-rays. upper panel: the first direct evidence of the pion bump from agile data below 200 mev (giuliani et al., 2011). lower panel: new fermi-lat data below 200 mev confirm the observed steepening due to the neutral pion emission (ackermann et al., 2013). preliminary results from a new detailed study of the snr gamma-cygni with agile have been presented by g. piano during the 11th agile workshop3 (piano et al., paper in progress). from agile data analysis there are hints of non-thermal emission possibly related to shock-cloud interactions in the north-western part of the shell. the agile minicalorimeter is also detecting terrestrial gamma-ray flashes (tgfs), intense and brief pulses of gamma-rays originating in the earth atmosphere, and associated with thunderstorm activity. tgfs last a few thousandths of a second, and produce gamma-ray flashes up to 100 mev, on timescales as low as < 5 ms (marisaldi et al., 2010a). agile joins other satellites in detecting tgfs, but its unique capa3http://www.asdc.asi.it/11thagilemeeting/program.php 160 agile highlights after six years in orbit bility of detecting photons of the highest energies within the shortest timescales makes it an ideal instrument to study these impulsive phenomena. agile data have shown for the first time that tgf cumulative spectrum at high energy deviates from a power law with exponential cutoff model and can be better fit with a broken power law with significant counts above background up to 100 mev. the crucial agile contribution to tgf science is thus the discovery that the tgf spectrum extends well above 40 mev, and that the high energy tail of the tgf spectrum is harder than expected. combined observations of high-energy photons detected by the agile-grid detector have also provided the first direct localization of tgfs in gamma-rays, within a region clustered around the sub-satellite point (fuschino et al., 2011). thanks to its very low inclination orbit in the ±2.5 latitude band, agile provides the largest tgf surface density across the equator available up to now. as statistics grows, agile data show that the tgf sample is possibly composed by two distinct populations: a low-energy population, consistent with typical tgf characteristics, and a high-energy population, with quite distinct features (tavani et al., 2011; marisaldi et al., 2012). crab nebula variability: the surprising discovery by agile of variable gamma-ray emission above 100 mev from the crab nebula in sept. 2010 (tavani et al., 2010; tavani et al., 2011b), and the fermilat confirmation (buehler et al., 2010; abdo et al., 2011) started a new era of investigation of the crab system. the 2012 bruno rossi international prize has been awarded to the pi, marco tavani, and the agile team for this important and unexpected discovery. astronomers have long believed the crab to be an almost ideal standard candle, a nearly constant source (at a level of few percent) from optical to gamma-ray energies (meyer et al., 2010), with possible long-term nebular flux variations over a few-year timescale reported in the hard x-ray range. on september 2010 agile detected a rapid gamma-ray flare over a daily timescale, see fig.4, and thanks to its rapid alert system, made the first public announcement on september 22, 2010. this finding was confirmed the next day by the fermi observatory. agile had also previoulsy detected a giant flare from the crab in october, 2007 during the initial science verification period of the satellite, and in the first agile catalog paper (pittori et al., 2009), it was reported that anomalous episodic flux values observed from the crab in 2007 were under investigation. we know now of five major gamma-ray flares from the crab nebula detected by the agile-grid and fermi-lat, up to may 2013. gamma-ray data provide evidence for particle acceleration mechanisms in nebular shock regions more efficient than previously expected from current theoretical models. we estimate a recurrence rate for strong gamma-ray flares of ∼ (1 − 2)/year. figure 5: the crab nebula flare in september 2010, as observed by agile at energies above 100 mev, (tavani et al., 2011b). a recent study (striani et al., 2013 ) of the flux and spectral variability of the crab nebula above 100 mev on different timescales, ranging from days to weeks, both in agile-grid and in fermi-lat data, also gives evidence for week-long and less intense episodes of enhanced gamma-ray emission that we call “waves” which can occur by themselves or in association with shorter flares. both the flaring and “wave” events can be related to a population of accelerated electrons consistent with a mono-chromatic or relativistic maxwellian distribution. gamma-ray observations challenge standard magneto-hydrodynamics models of nebular emission. 5 conclusions the study of cosmic gamma-rays in the energy range from a few tens of mev to a few tens of gev is only possible from space due to atmospheric absorption. with the successful launch of the new generation gamma-ray space asi telescope agile on april 2007, followed by nasa fermi on june 2008, gamma-ray astrophysics entered a new era. agile is substantially improving our knowledge on various known gamma-rays sources, such as supernova remnants and black hole binaries, pulsars and pulsar wind nebulae, blazars and gamma ray bursts. moreover, agile has contributed to the discovery and study of new galactic gamma-ray source classes, and of galactic gamma-ray transients. the mission is also giving a crucial contribution to the study of the terrestrial gamma-ray flashes seen in the earth atmosphere. gamma-ray observations challenge standard theoretical models of particle acceleration and dynamics under extreme matter conditions. acknowledgement research partially supported through the asi grants i/089/06/2, i/042/10/0 and i/028/12/0. 161 carlotta pittori references [1] abdo, a. a., et al. 2009, science, 326, 1512. doi:10.1126/science.1182174 [2] abdo, a. a., et al. 2011, science, 331, 739. doi:10.1126/science.1199705 [3] ackermann, m. et al., 2013, science 339, 807. 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[32] tavani, m. et al., 2011b, science, 331, 73. doi:10.1126/science.1200083 [33] verrecchia, f., et al., 2013, a&a 558, a137. discussion luigi piro: how can you distinguish between inverse compton and π0 decay in the gev range (question to manami sasaki talk on snrs, tue 28 afternoon). carlotta pittori comment: my comment to luigi piro question is that, as i have shown in my talk, at least in some cases as w44, in which middle aged snr interacts with molecular clouds, the new infrared planck data are very much constraining, and the overall multiwaveleght picture seem to largely favour π0 decay and hadronic models over the leptonic ones. 162 http://dx.doi.org/10.1126/science.1182174 http://dx.doi.org/10.1126/science.1199705 http://dx.doi.org/10.1088/0004-637x/775/2/98 http://dx.doi.org/10.1029/2009ja014502 http://dx.doi.org/10.1126/science.1183844 http://dx.doi.org/10.1088/2041-8205/712/1/l10 http://dx.doi.org/10.1038/nature08578 http://dx.doi.org/10.1126/science.1200083 introduction the ground segment and the agile data center new agile catalogs and asdc interactive web tools agile science highlights conclusions acta polytechnica ctu proceedings doi:10.14311/app.2015.1.0051 acta polytechnica ctu proceedings 2:51–56, 2015 © czech technical university in prague, 2015 available online at http://ojs.cvut.cz/ojs/index.php/app a case for domain-independent deterministic multiagent planning michal štolba department of computer science, faculty of electrical engineering, czech technical university in prague, karlovo nám. 13, 121 35 praha 2, czech republic correspondence: stolba@agents.fel.cvut.cz abstract. the notion of planning using multiple agents has been around since the very beginning of planning itself. it has been approached from various viewpoints especially in the multiagent systems community. recently, domain-independent multiagent planning has gained more attention also in the automated planning community. in this paper, we shortly present the current state of the art, question some aspects of the research field and discuss the rising challenges. keywords: planning, multi-agent planning. 1. introduction we could trace the first mention of multiagent planning back nearly as far as strips itself – in 1980 nillson published together with konolige a paper titled “multiple-agent planning systems” [1], in which they presented a high-level extension of strips [2] towards multiple agents. since then, the topic has been active mainly in the multiagent systems community. one of the most cited works on distributed and multiagent planning is [3], which describes basics of possible coordination schemes for planning agents. there was also a large amount of work dealing with another facets of the coordination area, but usually without deeper study of the plan synthesis part (e.g., gpgp [4]) or requiring additional domain-specific knowledge (e.g., talplanner [5]). multiagent planning also often relates to planning models described by decentralized pomdps (dec-pomdps) [6], similarly as pomdps are used in single agent planning under uncertainty. the point of view on multiagent planning from the multiagent community is extensively summarized in [7, 8]. in the planning community, extensions of the pddl language intended to support multiagent planning were introduced as multiagent planning language (mapl) in [9] and as multiagent pddl (ma-pddl) in [10], but none of the extensions gained wide popularity, probably because of their complexity. in 2008, a well accepted paper on multiagent planning was published by brafman and domshlak at the icaps conference [11], which finally ignited more interest in the topic. the paper formally introduced a minimalistic extension of strips (thus forming ma-strips) and have shown that the complexity of multiagent planning is not directly exponentially dependent on the number of agents, but rather on the tree-width of their interaction graph and a minimal number of interactions needed to solve the problem. such results suggested that at least for loosely coupled problems (where the tree-width is low), the approach may be beneficial. in the following years, several multiagent planners were proposed [12–19], either using directly the ma-strips formalism, or some ad-hoc, but very similar one. rather different approach was taken by crosby et al. in [20, 21], where the agent decomposition was based not on actions, but on a variables in finite domain representation, and was used in a centralized planner, significantly improving performance over (single agent) sota mainly in well decomposable domains. a dedicated workshop distributed and multiagent planning (dmap) also took place at the icaps’13 and icaps’14 conferences. 2. ma-strips ma-strips is the most commonly used formalism for domain-independent deterministic cooperative multiagent planning. the domain independence means that the input of the planner consists of description of available operators, predicates, functions etc. describing the general mechanics of the world as well as the particular instance of the world represented by ground initial state and the particular problem represented by the ground goal facts that need to be achieved. most commonly the input is described using pddl language, where the general part is termed “domain” and is described in a domain file and the particular initial and goal configurations are termed “problem” and are described in a problem file. in the multiagent setting, additional information is typically needed to determine what are agents and which actions belong to which agent (this is not part of the ma-strips formalism). deterministic in this case relates to the action model, where each action has deterministic effects, although the multiagent planning algorithm itself may be non-deterministic in the sense that for the same input it gives different output and runs for different time this is caused by the inherent nondeterminism of communication and distributed computation. 51 http://dx.doi.org/10.14311/app.2015.1.0051 http://ojs.cvut.cz/ojs/index.php/app michal štolba acta polytechnica ctu proceedings by cooperative multiagent planning we understand a setting where a group of agents is attempting to find a joint plan, where each agent uses its own actions in order to reach some joint (and possibly some private) goals. the agents typically need to coordinate their actions and to cooperate in order to do so, although each agent is planning for itself and the agents may have some private knowledge which they do not want to reveal to other agents. the ma-strips formalism is very much the same as the original strips, except that the actions are disjunctively split (or factored) among the agents, resulting in the following definition of the multiagent planning problem: definition 1. multiagent planning problem is a quadruple π = 〈l,a,s0,sg〉, where l is a set of propositions, a is a set of agents α1, . . . ,α|a|, s0 is an initial state and sg is a set of goal states. an agent α = {a1, . . . ,an} is represented by a set of actions it can perform. a state s ⊆l is a set of atoms from a finite set of propositions l = {p1, . . . ,pm} which hold in s. an action is a tuple a = 〈pre(a), add(a), del(a)〉, where a is a unique action label and pre(a), add(a), del(a) respectively denote the sets of preconditions, add effects and delete effects of a from l. in the set of propositions l can further be identified subsets of public propositions known to all agents and α-internal (a.k.a. α-private) propositions known only to a specific agent α. such separation of the propositions can be derived from the problem itself (as proposed in [11]) , where a proposition is public iff it is used by any two actions of different agents and it is internal to the agent α (α-internal) iff it is used only in actions of agent α. alternatively, what propositions are public or private can be a part of the problem definition (used in [16, 17]). how the actions are partitioned among agents is not a part of the ma-strips definition and is implementation dependent, in most planners, the agents are defined as objects in pddl and an action is assigned to an agent if the agent object is part of the grounded parameters of the action. similarly to the separation of the propositions, the actions of a single agent can be seen as either public or internal. an action a ∈ α is internal iff it uses only α-internal propositions in its preconditions and effects. otherwise, the action is public, which in fact means, that the action interacts with some other agent, or agents. following the strips formalism, a solution for the multiagent planning problem is a set of sequential plans (sequence of actions, one plan for each agent), where some of the actions of different agents may be performed in parallel, although in most implementations the resulting plan is sequential (i.e. there are no parallel actions). formally, the solution is defined as follows: definition 2. a solution to multiagent planning problem ∏ is a multiagent plan p =〈 pα1, ...,pα|a| 〉 , where each pαi is a plan of agent αi, i.e. consisting of a sequence of actions in αi ∪{noop}. a multiagent plan p is valid, iff (1.) for each α,β ∈a, |pα| = |pβ|. the noop actions mean that the agent is idle in the respective timestep. (a) for each time-step tk where k ∈ 〈1, ..., |pα|〉, actions ai = pαi[tk] for all i ∈ 〈1, ..., |a|〉are not mutually exclusive. (b) all actions in time-step t1 are applicable in the initial state, for all time-steps tk where k ∈ 〈2, ..., |pα|〉, actions in tk are applicable in the state resulting from time-step tk−1 and the state resulting from time-step t|pα| is in the set of goal states sg. since all the plans are required to have the same length, the ma plan can be represented by a matrix, where each row consists of a plan for single agent and each column contains all actions in the respective time-step. it is clear that the ma-strips formalism is very simplistic and minimalistic extension and can be developed further to accommodate various aspects of the multiagent systems (i.e., finer separation of public and internal actions, as discussed for example in [22]). the philosophy of the formalism is to provide the smallest possible common ground and to enable the easiest possible migration of the planning algorithm to the multiagent setting. unlike in many multiagent systems, the mastrips formalism assumes the agents to be fully cooperative. from the multiagent perspective, this may seem as an oversimplification, but similarly to the simplicity of the ma-strips, we argue to first tackle the problems of the simplest possible multiagent extension – cooperative agents – and then gradually add complexities of the multiagent nature, such as selfish agents, negotiation and such. some attempts have already been taken, for example in [23, 24]. 3. is it any good? the first question that is asked after introduction of some new or nonstandard paradigm is – what is it good for? how does it help us? in this paper, we would like to ask such questions and try to come up with some answers. from the multiagent point of view, we can see significant benefits of introducing domainindependent planning into multiagent systems, namely reuse of the wide spectrum of planning systems and techniques from the area of classical planning. on the other hand, the drawback of ma-strips for the multiagent setting is that it requires cooperation among the agents. this can be mitigated in future by means such as mechanism design, as shown in [24]. 52 vol. 2/2015 a case for domain-independent deterministic multiagent planning in the following section, we will present some of the benefits we suggest that multiagent planning may bring to the area of planning in general. 3.1. improve scalability by decomposition although being largely improved in recent years, scalability is still an issue for domain-independent planning. multiagent planning can be seen as a method of factorization of the planning problems, which has been studied in works such as [25, 26]. it has been shown that factorization of the planning problem can be beneficial, especially in the problems, where it is possible to find large independent parts – loosely coupled problems. take for example the rovers domain with several rovers. the problem may be factored in such a way that the rovers do not need to coordinate, except for the last bit, where they use shared communication channel to communicate the result. in such a case, most of the planning can be done independently (even on different machines), however some coordination is needed here (it is not possible to simply run several classical planners in parallel). this is a case for multiagent planning. given a single agent problem (e.g. from the ipc benchmark set), it is not obvious what entities should be considered as agents and how the factorization should be done. in ma-strips, the first question is not answered, but the second one has some theoretical background. once the actions are assigned to agents, the factorization is based on the interactions among the agents’ actions, effectively meaning that it is based on the causal graphs of the problem (as in [11]), the factoring is one of the best known to date, as shown in [25]. still, how to best assign actions to agents is an open question. in the fmap planner [17, 27] formalism, the actions are assigned to agents by the domain designer, as is the separation of private and shared information. this gives the domain designer more freedom, but also more responsibility in that making more information public may increase the problem complexity (as the tree width of interaction graph rises) and making less information public may render the problem unsolvable (i.e. global solution may not exist). completely different approach to factorization was taken by crosby in [20, 21], where the factorization is automatic and based on the variables of a finite domain representation (fdr) of the planning problem. the method is, again, based on causal graph structures and tries to identify agents as sets of variables which are somehow self-contained and represent the agent’s internal state. variables which do not form an internal state of any agent are understood as environment and thus public variables. classification of actions is rather more complex than in ma-strips, but can be summarized as follows. actions, which interact only with the agent variables are assigned to the particular agent, other actions are left as public, which means that they can be used by any agent. this approach was not shown to increase scalability, but was shown to increase solution efficiency, especially in some domains. note, that crosby’s approach was implemented as a centralized, single-thread planner. various kinds of factoring, and understanding gained from the research of it, may be also used for other aspects of classical planning, such as search state pruning [28]. 3.2. multi-core / cluster computing planning on multi-core machines gained wide popularity and had a separate track at the ipc2011 competition. it is definitely beneficial to be able to scale the computation using multiple cores. but, if we want to scale the computation even further and use multiple machines, we are no longer able to use shared memory (as in parallel computation). in such case, approaches of distributed computing utilizing message passing need to take place, and multiagent planning aspires to be one of them (although obviously not the only one). in this scenario, multiagent planning represented by ma-strips can be seen as an equivalent for distributed planning with factorization fixed by the partitioning of actions to particular agents. such approach has been successfully tested against sota in multi-core planning in [14] as a parallel version of mad-a*, called map-a*. again, the question of best factorization arises here, with the crosby’s variable based factorization being another candidate, however never tested in a distributed system. 3.3. privacy in multiagent systems, privacy of the data is often one of the main concerns. in ma-strips-based multiagent planning, this requirement is not one of the top priorities – the privacy (or internality) of the propositions and actions is understood rather as a way of improving effectiveness of the computation by reducing the complexity – but nonetheless if required by the application, some degree of privacy of the data can be achieved. in practice, it may be somewhat harder not to reveal the structure of the private information and it may also reduce usability of some techniques, such as distributed heuristics, but it seems to be possible. one such approach was described in [22], but without implementation or experimental evaluation. 3.4. asynchronous computation distributed computation (and multiagent as well) is inherently asynchronous and non-deterministic. this brings several challenges and complexities that need to be tackled, but once solved, some of the insights may be beneficial to classical planning as well. as a representative example, we can take the computation of a global heuristic. a heuristic in multiagent planning can be computed locally, using only a projection of the planning problem to the propositions of the respective agent. a public 53 michal štolba acta polytechnica ctu proceedings action a ∈ α can be projected to agent β by removing all α-internal propositions from pre(a), add(a) and del(a), thus retaining only public propositions (and technically also β-internal propositions, but actions of agent α contain none). such heuristics were used for example in [14]. a different approach is to attempt to compute a global heuristic estimate without exposing the internal information to other agents. this can be achieved for example by requesting other agents for their estimates of the current state and the goal, or some sub-goals, as in [29, 30]. such global heuristic computation may involve nontrivial communication, resulting in an asynchronous computation of the heuristic estimate. to our knowledge, such a phenomenon has not been studied yet. it seems to be interesting also from the point of view of classical planning, because if successfully utilized, asynchronous computation of heuristics may be used to evaluate computationally intensive heuristics, to use external solvers for heuristic estimation, sensory input (in robotics), etc. also it is interesting to observe that the global heuristic is always equally or more informative than the projected heuristic, thus we have two heuristics, the first one is less informative, gives lower estimates, but is faster, the second one is more informative, gives higher estimates, but takes significantly longer to compute. if generalized, a sequence of gradually more precise, but costlier heuristics could be used to guide search (including blind search as the fastest one), which is an interesting research topic. also, as the costly distributed heuristic is computed partly by other agents, the time while waiting for the results may be used to do some more computation such as exploration using only the local heuristic. 3.5. applications ma-strips planning may bring domain independent planning to some (more or less) new application domains. as such we see the following: • multi-robotics: slowly but steadily, domainindependent planning is penetrating the robotic research, providing the robots with high-level reasoning. on the other hand, robotic research is being extended towards multi-robotic teams. multiagent planning seems to bridge those two advancements and in the future may enable multi-robotic teams to be controlled by a high-level distributed planning system. in such scenarios, each robot would have its own planning representation of the world, thus reducing the domain and problem size and keeping the local information (i.e., sensory data) local. communication will clearly be an issue, which may encourage research of multiagent planning techniques aiming not primarily on planning speed, but on the communication requirements (this may for example discourage the use of global heuristic estimates). • orchestration of internet services: since internet services are computer programs already described in (formal) programming languages their coordination and orchestration can straightforwardly profit from multiagent planning techniques. moreover services are usually described by interfaces resembling planning actions with preconditions on the input data and effect on the data. therefore if such actions have to be ordered in a suitable fashion, beginning with initial constraints on the data and with a goal form of the data, planning is an appropriate fit. additionally, the services can be distributed over the internet and the data processing tasks can vary in the sense of particular areas of focus (e.g., financial processing, logistics services, etc.). both of these challenges are covered by the domain-independent multiagent planning. a single agent planning applications of this kind were already covered for example in [31]. • coalition planning: in recent years, many military operations involve coalitions. coalition operations are in nature cooperative, but some information cannot be disclosed among the coalition partners. this can be generally modeled by multiagent systems and in particular by multiagent planning. such a scenario is not applicable only in military, but may find uses also in business cooperation and other areas, although some more detailed privaci concerns may be necessary as illustrated in [22]. 4. challenges in the following, we will present some of the outstanding challenges of domain-independent multiagent planning we are aware of. 4.1. comparison of planners up to the present day, we know of about a half dozen domain-independent multiagent planners [12–17] and more are in the development. the biggest issue in order to compare them is that although the underlying formalisms are mostly similar, the actual language used as input differs (it is pddl with some ad-hoc definition of agents and/or public propositions). some consensus on this matter needs to be settled. similarly to the common definition language, a widely accepted set of benchmarks is needed. in the recent works, the benchmarks were typically created by ad-hoc converting some suitable ipc domains, where the agent decomposition is natural, such as logistics or rovers. only few new – multiagent specific – domains were introduced, i.e., cooperative pathfinding. in the future, it would be interesting to come up with more such domains, drawing from real world and multiagent systems applications and also designed to test some specific properties and pitfalls of multiagent planning. naturally, the best venue for such efforts would be a dedicated ipc track. the metrics used for planner comparison are also challenging. all metrics used to compare classical 54 vol. 2/2015 a case for domain-independent deterministic multiagent planning planners still apply, but there are some additional ones, such as number of exchanged messages or the amount of communicated data. further metrics can be adopted from the research of other multiagent algorithms. moreover, the comparison of multiagent planners gains a new complexity in that a single problem can be partitioned in multiple ways, and each partitioning may be beneficial for different planners. similarly, the amount of agent interaction significantly influences performance of the planners some algorithms may be better suited for loosely coupled problems while other for tightly coupled problems. this should be also reflected in the potential analysis. 4.2. which planning paradigm is the best? in classical sequential domain-independent planning, the dominating planning paradigm seems to be heuristic forward-chaining search typically utilizing one or more highly informative heuristic estimators (used by planners as fastforward, fastdownward, lama and many more), although some other approaches (such as bidirectional search) has performed well in the last ipc competition. it is not clear, whether the same holds for multiagent planning (and for all metrics). again, the only way to reliably answer this question is to perform a rigorous comparison of the planners, such as in the ipc competition. currently the most frequent approaches to multiagent planning are distributed heuristic search [14, 29, 30], partial order planning [16, 17], and various other, such as plan reuse [12] or generate-and-test [18, 19]. 4.3. heuristics one interesting area of research are the heuristic estimator for multiagent planning. as mentioned in the section about asynchronous computation, the baseline solution for heuristic computation is to use a projection of the problem to the particular agent’s propositions. this approach was used in [13, 14]. another approach is to attempt to compute (or at least approximate) the global value as if the problem was not partitioned in order to obtain better heuristic guidance. to our knowledge, the only heuristics treated this way so far were the ff heuristic in [29] and [30] and a domain transition graph based heuristic in [17]. of course, in some cases, the communication needed to compute a global heuristic may be prohibitive, in such a situation, the local heuristic would have to suffice, or a clever combination of both would need to be devised. different challenge is to come up with a heuristic specific to the multiagent planning. such heuristic may not only lead the search towards the goal, but may also lead the search in a way minimizing some multiagent metric, such as the communication load. 4.4. privacy multiagent planning can be seen also from the perspective of privacy-preserving distributed computing, although it is not the main concern of most of the ma-strips based planners. this perspective raises many questions, some of them were already addressed in [22, 32], but a comprehensive analysis of possible privacy violations in ma-strips and ma-strips based planning algorithms has not been yet presented. even the question if a ma-strips based multiagent planner can be privacy preserving (and if so how strongly) is yet to be answered. 4.5. extending complexity analysis from the theoretical perspective, the original [11]’s complexity results can give a solid groundwork for future and more detailed studies of complexity of multiagent planning. we envision two main directions: (i) problem-specific and (ii) finer complexity classes. the problem-specific study could be inspired by works as [33], where particular planning domains are analyzed from perspective of computational complexity showing which problems are easy, even if planning in general is known to be computationally hard. the idea of study of finer complexity classes, is for example parametrized analysis of classical planning by [34]. additionally, these studies can be extended by communication complexity showing how much communication would be required to solve the problems. 5. final remarks domain independent planning is a field of research bridging two distinct topics of multiagent systems and domain independent planning. we see providing a common problem definition language and set of benchmarks in order to be able to determine the state of the art and compare it with newly emerging techniques 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[34] c. bäckström, y. chen, p. jonsson, et al. the complexity of planning revisited a parameterized analysis. in aaai. 2012. 56 http://dx.doi.org/10.1609/aimag.v22i3.1581 http://dx.doi.org/10.1613/jair.1497 http://dx.doi.org/10.3233/mgs-2009-0133 http://dx.doi.org/10.1145/1838206.1838379 http://dx.doi.org/10.5220/0002702601280134 http://dx.doi.org/10.3233/978-1-61499-098-7-762 http://dx.doi.org/10.5220/0004918701780189 http://dx.doi.org/10.1007/s10489-014-0540-2 http://dx.doi.org/10.3233/978-1-61499-098-7-624 acta polytechnica ctu proceedings 2:51–56, 2015 1 introduction 2 ma-strips 3 is it any good? 3.1 improve scalability by decomposition 3.2 multi-core / cluster computing 3.3 privacy 3.4 asynchronous computation 3.5 applications 4 challenges 4.1 comparison of planners 4.2 which planning paradigm is the best? 4.3 heuristics 4.4 privacy 4.5 extending complexity analysis 5 final remarks references acta polytechnica ctu proceedings doi:10.14311/app.2016.3.0056 acta polytechnica ctu proceedings 3:56–59, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app determination of thermal response of carrara and sneznikovsky marble used as building material veronika petráňováa, ∗, jaroslav valacha, alberto vianib, marta peréz estébanezb a institute of theoretical and applied mechanics as cr, v.v.i., prosecka 76, prague, czech republic b institute of theoretical and applied mechanics as cr, v.v.i., centre of excellence teläŋ, batelovska 485, teläŋ, czech republic ∗ corresponding author: petranova@itam.cas.cz abstract. physical weathering of marble, widely used as a cladding material on buildings, is one of the most common damaging mechanism caused by anisotropic thermal expansion of calcite grains. the extent of marble deterioration depends mainly on stone fabric and texture. dry cuboids of carrara marble and marble from dolni morava quarry were subjected to microscopic analysis and thermal cycling, to determine the thermal expansion related to stone fabric and predominant lattice orientation of grains (i.e. texture). keywords: marble, anisotropic thermal expansion, thermal dilatation. 1. introduction marble has been predetermined as building and decorative natural stone since antiquity because of its good workability and unique visual appearance. these properties are related to the composition of carbonate sediments and to the nature and intensity of the recrystallization process under the regional metamorphosis conditions. marble is considerably vulnerable to both chemical and physical weathering by reason that this type of stone predominantly consists of carbonate minerals calcite caco3 and dolomite camg(co3)2. periodic temperature variations, together with changing in moisture can cause serious damage to the stone structure, finally leading to the complete stone decay. the enormous sensitivity of marble to variations in temperature originates in the thermal response of calcite crystals. while the dolomite crystals subjected to thermal cycling expand only, the calcite crystals show anisotropic thermal dilatation introduced by thermal cycles, i.e., it expands along c axis during heating and contracts in other directions. however, cooling of calcite crystals leads to the opposite effect, i.e., contraction along the c axis and expansion in the other directions as is illustrated in figure 1. both mechanisms are considered as the major cause of creation and propagation of cracks along the grain boundaries in the stone. thermal expansion can be expressed using the thermal expansion coefficient α, describing the length change per unit of temperature. the coefficient value depends on the considered temperature interval. to quantify the extent of stone elongation, the residual strain ε is more suitable [1]. the preferred lattice orientation of calcite grains i.e. texture is one of the most important stone characteristics, essentially influencing thermal response in different spatial directions. marble with strong texture is generally more vulnerable to thermal weathering, which can be enhanced by significant differences between the α values along the cand a-axis of individual calcite grains [2]. thermal expansion of marble is measurable at low temperatures enabling simulation the climate conditions in the central europe. nevertheless, experimentally obtained thermal dilatation values often deviated from calculation based on theoretical models considering only the texture variations in a marble. another important parameters are grain shapes, grain sizes and the type of grain boundaries because of the strong influence on the rate of stone decay. a certain directional relationship between the shape preferred orientation (spo) of calcite grains and their lattice preferred orientation (lpo) can extensively intensify the thermal response of marble. for example, marble with lpo parallel to spo exhibits a strong thermomechanical response because of the large concentration of maximum principal stress in the direction of shape elongation. thermal stresses during heating can initiate the formation of new cracks or widening of preexisting cracks or pores. microcracking formation can also be influenced by grain size since the level of marble decay is also connected with the number of adjacent grains and with the grain boundaries [3–5]. however, thermal response and consequently the rate of physical deterioration of marble depends also on other stone characteristics, including the ratio of calcite and dolomite, moisture content, porosity and pore size distribution of the investigated marble [6]. depending on all the mentioned stone characteristics, the thermal response of marble after cooling down to room temperature can be determined as isotropic or anisotropic, with or without residual strain [1]. 56 http://dx.doi.org/10.14311/app.2016.3.0056 http://ojs.cvut.cz/ojs/index.php/app vol. 3/2016 thermal response of carrara and sneznikovsky marbles figure 1. anisotropic thermal expansion of calcite crystals. adopted from [1]. 2. materials and methods testing of the thermal response was performed on two structurally different marble types. the first one from the carrara area in northern italy was chosen for its simple fabric. the second one sneznikovsky marble was taken from dolni morava quarry in the czech republic. the texture as well as the shape and size of grains was investigated by optical microscopy and electron back-scattered diffraction (ebsd) method in a scanning electron microscope (sem, fei quanta 450, fei corp., usa). for the sem-ebsd investigations, slide specimens of marble were coated with a thin layer of carbon to protect against specimen charging. to evaluate the dependence of dilatation on the sample’s microstructure, two pairs of perpendicular cuboids a-a, b-a, c-1 and c-4 (6 x 6 x 20 mm) were prepared from the original marble blocks. the orientation of the cuboid samples is depicted in figure 2. thermal expansion measurements were performed using a vertical dilatometer (l75 pt linseis corp., germany). the specimen elongation was measured after each 1 °c step. the cuboid samples were subjected to heating and cooling cycles, ranging in temperature from -40 to 100 °c. the cycles were repeated 7 times for specimens a-a, b-a and c-1 and 5 times for specimen c-4. the temperature interval was gradually increased for the a-a, b-a and c-1 specimens. the interval for specimen c-4 remained stable in all 5 cycles (from -40 to 100 °c). the values for the thermal expansion coefficient α values were calculated using equation (1). the length change in 4l was determined within a temperature interval 4t, from 18 °c to the maximum temperature in the considered cycle, i.e. for cycle 1, the maximum temperature was 33 °c, whereas for cycle 7, it was 100 °c. to determine α, only a positive temperature range was selected, because the results from the range below the freezing point were not reliable. α = 4l l4t (1) figure 2. orientation of cuboid specimens, a-a and b-a is sneznikovsky marble and c-1 and c-4 is carrara marble. the symbols and their physical meanings are listed at the end of the paper. 3. experimental results based on optical and sem-ebsd observations, the carrara marble exhibits a weak texture in both directions (x-y and x-z). specimens c-1 and c-4 contain calcite grains with a nearly uniform size of around 0.3 mm with straight polygonal grain boundaries and no shape preferred orientation. contrary to carrara, sneznikovsky marble exhibits cracking, predominantly on planes parallel to the x-y plane. the microfabric of 57 v. petráňová, j. valach, a. viani, m. peréz estébanez acta polytechnica ctu proceedings figure 3. a microfabric of investigated samples, b orientation of calcite grains in b-a sample. figure 4. dependence of relative extension on the temperature. the investigated marble types is illustrated in figure 3. the size of the calcite grains in the a-a and b-a specimens is not uniform; grains in sizes up to 4 mm with sutured grain boundaries can be found. as in the carrara specimen, both perpendicular planes of sneznikovsky marble contain grains with no shape preferred orientation. the elongation of the cuboid specimens was expressed as relative extension, i.e. percentage difference between 4t for each temperature step and original length of the specimen. the highest relative extension of 0.18% was observed in c-4 at 100 °c in the first cycle. in the other cycles (2-5) the value was equal to 0.16%. specimen c-1 showed an increasing trend over the temperature and reached its maximum relative extension of 0.14% at 100 °c. regarding the coefficient α, specimen c-4 also showed the highest average value of 12.5 · 10−6k−1. specimen c-1 achieved the most significant difference in α, varying from 3.9 ·10−6k−1 to 11.3 · 10−6k−1. the results are illustrated in the figures 4 and 5. the first cycle of the measurements performed on the specimen c-1 was excluded from the results. values obtained from the first cycle performed in c-1 specimen weren’t included to the results. during the 7 cycles in specimens a-a and b-a relative elongation also showed an increasing trend, but the maximum values were more different than in the carrara marble. at 100 °c in cycle 7, the relative figure 5. thermal expansion coefficient α of investigated specimens, loaded in the 18-100 °c temperature interval. the value marked "a" represents an average value of α determined by [7] on large set of marble specimens. extension in specimen a-a was equal to 0.15% while in b-a it was just 0.10%. the α coefficients determined for the sneznikovsky marble were not as variable as in the carrara type. in the a-a sample, α reached an average value of 10 · 10−6k−1 and in the b-a sample of 5 ·10−6k−1. based on the relative extension difference between corresponding cycles, an α variation of 5 ·10−6k−1 was determined, which can be attributed to the observed anisotropy in the sneznikovsky mar58 vol. 3/2016 thermal response of carrara and sneznikovsky marbles ble, i.e. the x-y orientated plane in specimen a-a was more fissured. 4. conclusions it can be concluded that simultaneous microscopic and dilatometry investigations can help in elucidation of thermally induced degradation of marbles. based on the results, the thermal response and generation of damage caused by accumulation of residual strains related to stone fabric and texture in investigated marbles will be determined on a larger set of perpendicular specimens, heated and cooled in the positive range of temperatures (i.e. from 20 to 100 °c). investigations carried out on both specimen types reveal an increasing rate of thermal expansion with increasing temperatures. in presence of a temperature gradient, this characteristic can contribute to the well known bowing effect on marble slabs. list of symbols α thermal expansion coefficient [ k−1] 4l difference between the actual specimen length and the length at the room temperature [m] l length of the specimen at the room temperature [m] 4t temperature difference between the actual and the room temperature [k] acknowledgements the research has been supported by czech science foundation (project no. p105/12/g059). references [1] m. steiger, a. charola. weathering and deterioration. in the stone in architecture, chap. 4., springer-verlag, 2011. doi:10.1007/978-3-642-14475-2_4 [2] s. siegesmund, k. ullemeyer, t. weiss, e.k. tschegg. physical weathering of marbles caused by anisotropic thermal expansion. inter j earth sci 89(1):170-182, 2000. [3] v. shushakova, e.r. fuller jr, s. siegesmund. influence of shape fabric and crystal texture on marble degradation phenomena:simulations. env earth sci 63(7-8):1587-1601, 2011. doi:10.1007/s12665-010-0744-7 [4] b. leiss, t. weiss. fabric anisotropy and its influence on physical weathering of different tyes of carrara marbles. j struct geol 22(11-12):1737-1745, 2000. [5] u. akesson, j.e. lindqvist, b. schouenborg, b. grelk. relationship between microstructure and bowing properties of calcite marble claddings. bull eng geol env 65(1):73-79, 2006. [6] e. cantisani, e. pecchioni, f. fratini, c.a. garzonio, p. malesani, g. molli. thermal stress in the apuan marbles: relationship between microstructure and petrophysical properties characteristics. inter j rock mech min sci 46(1):128-137, 2009. doi:10.1016/j.ijrmms.2008.06.005 [7] a. zeisig, s. siegesmund, t. weiss. thermal expansion and its control on the durability of marbles. in the natural stone, weathering phenomena, conservation strategies and case studies, chap. 2., the geological society of london, 2002. 59 http://dx.doi.org/10.1007/978-3-642-14475-2_4 http://dx.doi.org/10.1007/s12665-010-0744-7 http://dx.doi.org/10.1016/j.ijrmms.2008.06.005 acta polytechnica ctu proceedings 3:56–59, 2016 1 introduction 2 materials and methods 3 experimental results 4 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2015.1.0008 acta polytechnica ctu proceedings 2:8–14, 2015 © czech technical university in prague, 2015 available online at http://ojs.cvut.cz/ojs/index.php/app on fpga based acceleration of image processing in mobile robotics petr čížek∗, jan faigl department of computer science, faculty of electrical engineering, ctu in prague, technická 2, 166 27 prague, czech republic ∗ corresponding author: petr.cizek@fel.cvut.cz abstract. in visual navigation tasks, a lack of the computational resources is one of the main limitations of micro robotic platforms to be deployed in autonomous missions. it is because the most of nowadays techniques of visual navigation relies on a detection of salient points that is computationally very demanding. in this paper, an fpga assisted acceleration of image processing is considered to overcome limitations of computational resources available on-board and to enable high processing speeds while it may lower the power consumption of the system. the paper reports on performance evaluation of the cpu–based and fpga–based implementations of a visual teach-and-repeat navigation system based on detection and tracking of the fast image salient points. the results indicate that even a computationally efficient fast algorithm can benefit from a parallel (low–cost) fpga–based implementation that has a competitive processing time but more importantly it is a more power efficient. keywords: fpga, system–on–chip, image processing, fast feature detector, visual navigation. 1. introduction one of the key indicators of the intelligent behaviour of a mobile robot is its ability to react promptly in complex situations and make fast and correct decisions regarding the robot goals. on micro-robotic platforms, this is a very comprehensive task due to limited computational resources. a typical example of micro-robotic platforms are micro aerial vehicles (mavs) [1], small legged robots [2] or robots used to study swarm intelligence [3]. they can be characterized by constrained dimensions and payload, which is directly related to the limited battery capacity. altogether, these parameters determine an available computational power on-board of the mobile robot. a direct trade-off between the computational capabilities and power consumption can be identified, characterizing the maximum time for which the robot can perform its mission. this trade-off between computationally demanding and power efficient decision-making techniques is especially recognizable in navigation algorithms that rely on computer vision methods based on processing of a large amount of visual data. salient point extractors are popular techniques of machine perception in the mobile robot navigation tasks; however, not all available approaches are currently suitable for micro-robotic platforms due to the limited on-board computational resources. moreover, a practical deployment of a mobile robot always demands a real-time performance of decision-making algorithms, which imposes further restrictions on the applicable approaches regarding the available resources. two fundamental approaches how to deal with limited computational resources on micro-robotic platforms can be considered. the first approach is a simplification of the computationally demanding processing, e.g., by introducing approximations of the demanding method and simplifying the whole principle, which can be a daunting task as it may not be always possible. the second approach is a utilization of the optimized implementations for newly available features of the modern processors, e.g., a dedicated optimization for special (typically simd– single instruction multiple data) instructions of the new cpus. besides, dedicated co-processors can be utilized to speedup the computationally demanding calculations by massively parallel processing that is available using conventional graphics cards, e.g., using cuda or opencl. although modern graphics cards are able to provide superior peak performance, they also have high power consumption requirements (in order of hundreds of watts) and thus they are not suitable for micro-robotic platforms. therefore, as an alternative solution for parallel based and power-efficient computations, it is a more suitable to consider field-programmable gate array (fpga) technology to develop a custom architecture specifically designed for the particular computational task, which, in the end, can be very power and computationally efficient while also small in dimensions and cheap in development. however, fpga-based solutions need a significantly different approach for an efficient implementation of the particular algorithms than a conventional cpu. thus, the deployment time can be high compared to a gain from the implementation of a custom design computational architecture for the fpga. in this paper, we report our results on a comparison 8 http://dx.doi.org/10.14311/app.2015.1.0008 http://ojs.cvut.cz/ojs/index.php/app vol. 2/2015 on fpga based acceleration of image processing in mobile robotics of the selected image processing techniques suitable for embedded computation of the visual navigation task implemented on a conventional on-board platform and a dedicated solution based on fpga processing of the image data. in particular, we consider a detection of the salient points in the image. the salient points are image patterns which differs from their local neighbourhood and are expected to be reliably and repeatably detectable, preferably invariant to camera viewpoint changes. such features provide a mobile robot with a limited set of environmental anchor points which are then utilized for a vision-based navigation. feature based methods consist of three stages. the first stage is a detection of features to identify salient points in the image. then, for each detected feature its descriptor is calculated to describe the local image surrounding of the feature in order to distinguish individual features. the third stage is a process of establishing feature correspondences which is called feature matching. the matching is based on a comparison of the feature descriptors to determine whether the features correspond to the same salient object already detected in the environment. regarding the computational complexity of the feature detection, we can classify the features to artificial, like blobs or patterns, and naturally occurring in the environment. for the field robotics, natural landmarks are more important; however, artificial landmarks are much easier to detect. one of the foundational feature detectors is the scale-invariant feature transform (sift) algorithm [4], which is unfortunately computationally very demanding, and therefore, a speeded-up robust features (surf) algorithm has been proposed [5] to provide salient points with less computational requirements. although surf is less demanding than sift, it can be still computationally restrictive on small platforms. that is why researchers investigate other methods how to detect and describe salient objects in the environment by a computationally efficient algorithm such as the fast feature detector [6]. fast is widely adopted by robotic community for its low computational complexity. it has been deployed in several visual navigation methods, e.g., [7, 8], and it also provides a base for different feature extractors [9, 10]. following our intention to have a power and computationally efficient image processing available on a mobile robot, we consider fast as a suitable algorithm for visual navigation task computed on-board. therefore, we compare a performance of the fast feature extraction implemented on a conventional embedded platform with a cpu based on an arm architecture and a custom implementation of the fast detector on an fpga-based dedicated co-processor. the comparison indicates that even a computationally efficient feature extraction based on the fast algorithm can benefit from parallel fpga-based implementation that has a competitive processing time figure 1. hexapod walking robot platform and it is a more power efficient. the paper is organized as follows. an overview of the most related reports on enhanced computationally and power efficient vision-based navigation approaches based on fpga co-processor is presented in section 2. the proposed evaluation of the fpga-based image processing is considered in the context of a monocular vision-based teach-and-repeat autonomous navigation [11] that is briefly described in section 4 together with the main idea of the fast feature detection and the selected brief feature descriptor [12]. results on the evaluation of the computational burden reduction using an embedded platform of the hexapod walking robot (see figure 1) and fpga–based solution are presented in section 5. concluding remarks and future work is dedicated to conclusion in section 6. 2. related work this paper is focused on fpga-based implementations of computer vision techniques that are utilized in mobile robot navigation tasks. several approaches have been published and their authors reported improvements of the computational and power efficiency in vision-based navigation tasks by a dedicated implementation on fpga. probably one of the first fpga based implementation of the surf feature extraction has been introduced in [13], where an embedded module capable of image processing suitable for mobile robotics is presented. the module implements a surf feature extraction [14] from images with the resolution of 1024×768 pixels. the reported frame rate is 15 fps, which can be considered low in a comparison to the 24 fps gpu-based implementation of surf; however, the power consumption is only 6 watts. notice, the fpga implementation outperforms a pure cpu implementation running on an intel atom dual-core processor, which is able to process only a single frame per second. another fpga-based module for stereo image processing has been recently introduced by the eth 9 petr čížek, jan faigl acta polytechnica ctu proceedings computer vision and geometry group [15]. it is based on the fpga and arm cortex a9 quad-core cpu and calculates dense disparity images with the resolution of 752×480 pixels at the 60 frames per second. the disparity estimation relies on the semi global matching approach, which is computationally a very intensive task. the reported power consumption of the module is 5 watts. in [16], the same authors extend their module by implementing a reactive based collision avoidance method for mavs and tested it in an outdoor environment. authors of [17] presented a miniature module with a low-cost fpga and mcu for visual navigation of mavs. they utilize a reduced version of the ptam algorithm [8] (mapping maximum of 200 features due to memory limitations) for an estimation of the visual odometry. their approach is based on the fpga implementation of the fast features detection and brief [12] feature description. the mcu is utilized for visual odometry calculation and the whole system operates at 30 frames per second on images with the resolution of 160×120 pixels. in [18], authors consider the fpga to synchronize data from the imu and fast feature extraction from a camera stereo pair for a robust inertial assisted realtime visual slam. the reported update rate is about 20 frames per second for images with the resolution of 752×480 pixels. similar results with a slightly different approach are presented in [19]. an embedded fpga–based computer vision module has been presented in [20]. the proposed fpga architecture uses an efficient pipelining for the fast feature detection in a video stream with the resolution of 752×480 pixels at the frame rate of 60 frames per second, which is utilized in the visual teach–and– repeat navigation method. regarding the physical dimensions of the aforementioned modules, they are all based on system–on–chip (soc) solutions on fpga development boards which (in all cases) do not exceed 12 cm × 8 cm. the power consumption of these modules is less than 10 watts while all of them exhibit a real-time performance in the particular robotic navigation task. the soc architecture combines the advantages of the fpga with a regular cpu in order to reduce deployment costs and improve performance of the system. its main aspects are described in section 3. all of the discussed approaches seem to be a suitable choice for deployment on a micro-robotic platform. however, the most promising are the approaches ([17– 20]) based on the fast feature detector [6] because this detector is widely adopted by the robotic community due to its low computational complexity. hence, it represents a base for further methods. for example, it is a source of the parallel tracking and mapping (ptam) method [8] that can be considered as a visual monocular odometry (or slam – simultaneous localization and mapping) technique, also suitable for mavs, albeit the original ptam method was designed to be used in small augmented reality workspaces. the method relies on tracking and mapping of fast features for the 6dof position estimation in the environment. fast has been recently utilized in the semi-direct monocular visual odometry [7] in which fast features are tracked and used to build a map that is further used for a visual odometry estimation. the algorithm was tested on mav equipped with the odroid-u2 arm quad-core computer achieving computational speed of 50 frames per second. despite of the cpu implementation, the reported power consumption of the system is 10 watts, which is a competitive to the one of the first fpga module for the surf detection introduced in [13]. therefore, we consider the fast features detection on the arm–based computer and fpga–based implementation [20] in this evaluation study of the potential benefits of the parallel (fpga–based) implementation of the image processing for visual navigation. 3. processor centric design a combination of the fpga and cpu allows to accelerate a computer vision application and also allows to carry out more complex tasks by an improved system. the main idea of combining these two types of computational architectures is the processor centric system–on–chip co-design, which allows to exploit benefits of both architectures. generally, a cpu is a more suitable for general purpose computations and it is also a more easily programmable. however, a parallel nature of the fpga fabric makes it well suitable for accelerating demanding or repetitive computational operations performed by the cpu. besides, it is also suitable for online signal processing and sensory interfacing [21]. there are two principal ways how to incorporate both the fpga and cpu in a single embedded design. it is possible to have both of them soldered on the printed circuit board and utilize a communication channel like spi for a fast data transmission between them. however, much more efficient is to utilize the system-on-chip (soc) solution, where the cpu is a part of the fpga chip. the cpu can be either softcore or hardcore. the fpga fabric is extremely versatile, and therefore, it can implement even a cpu in its fabric, which is called a softcore cpu. on the other hand, manufacturers started to incorporate the cpu in the design of the chip to further exploit the fpga–cpu co-design possibilities. in that case, the cpu is etched into the silicon of the chip together with the fpga fabric and such a solution is then referenced as the hardcore cpu. nowadays hardcore processors are usually based on multi-core arm architecture and they are running at the units of ghz. in the case of the softcore processors, their speed is limited to around 200 mhz due to the fpga fabric limitations. the main advantage of the both hardcore and softcore processors is that they 10 vol. 2/2015 on fpga based acceleration of image processing in mobile robotics can be directly programmed in the c programming language. in addition, they can host a real-time operating system, which makes their programming for time critical applications even simpler. from the processor point of view, the fpga fabric can act either as a memory mapped slave device or it can be directly connected to the cpu core and accelerates the processor instructions. in that case, the instruction set of the cpu is extended by custom designed instructions. the main advantage of fpga–based soc systems is their versatility and reconfigurability which may be of a particular use in the field of mobile robotics. the programmer can choose components which assemble the system according to the needs of the particular application. the fpga resources allow to build both internal computational units and io communication interfaces, which can be directly connected to the main cpu. moreover, fpga always posses a lot of general purpose io pins which can be used for any communication not exceeding maximum frequency of the given fpga chip (usually about 300∼350 mhz) regardless the communication, which can be serial like spi and i2c or parallel. in addition, specialized hardcore bus endpoints for a high-speed communication are common integral parts of nowadays fpgas, e.g., gigahertz serial endpoints like pcie, sata or parallel dram interfaces. 4. bearing-only visual navigation the navigation algorithm used for the evaluation of implementations of the fast feature detector on the cpu and fpga is based on the teach-and-repeat navigation algorithm proposed in [11]. the navigation is based on tracking previously mapped visual features that are used for a bearing-only navigation. the approach relies on the relative localization provided by the odometry, which would eventually integrate unbounded position error over time. however, the approach considers the odometry only locally for traversing a short straight line segment along which detected visual landmarks are utilized to correct the robot heading and thus suppresses the odometric error. the corrections of the heading are based on a modus of the horizontal displacements of the tentative correspondences between previously mapped and the currently perceived visual features. the modus is found by a histogram voting as it is visualized in figure 2. the tentative correspondences are established using the fast feature detector and the brief feature descriptor [12] which are briefly described in the following subsections. 4.1. features from accelerated segment test the fast feature detector [6] belongs to the family of the so called appearance based detectors because it figure 2. surfnav navigation algorithm. matched features from the current and previously learned image. the corresponding navigation histogram is depicted at the bottom. searches the image for corner-like structures. the detector is optimized with respect to the computational complexity. the algorithm uses a set of comparisons between the center pixel p and pixels defined on the 7×7 neighbourhood using the circle in the image determined by bresenham’s circle algorithm, which is shown in figure 3. figure 3. fast feature detection. courtesy of [6]. at first, the pixels on the circle are labelled dark or bright depending on their relative brightness to the central pixel. the candidate pixel is considered as a feature if there are n contiguous bright or dark pixels in the circle. for n ≥ 12 it is possible to use a rapid rejection method for a faster outlier rejection which leads to a common setting of n = 12. however, the original paper [6] approved that the best repeatability is exhibited for the setting of n = 9. the corner score, which is necessary for the non-maxima suppression, is calculated as the sum of the absolute differences between the pixels in the contiguous arc and the central pixel. the fpga architecture presented in [20] uses efficient pipelining for the fast feature detection acceleration. the image data are processed as an online stream and the architecture performs all the steps of the algorithm, i.e., the corner detection, corner score calculation, and non-maxima suppression, in parallel as data are transmitted from the image sensor. the architecture also allows to use different values of n without an impact on the computational speed or used 11 petr čížek, jan faigl acta polytechnica ctu proceedings resources. 4.2. binary robust independent elementary features once features are detected, individual salient points are characterized by descriptors. the purpose of the descriptor is to describe the image neighbourhood of the point and thus characterize the salient object of the environment. in the proposed evaluation, the selected descriptor is the brief descriptor which stands to the binary feature descriptor [12]. it is based on pairwise intensity comparisons of pixels inside an image patch surrounding the located feature. these comparisons form a set of unique binary tests which are subsequently stored into a q-dimensional bit vector. the pairwise comparisons can be chosen either randomly or evolutionary, e.g., they can be trained by methods of reinforcement learning [22], which optimize the descriptor for the particular environment. typically used values of q are 64, 128, and 256 bits that correspond to the brief8, brief16, and brief32 variants of the feature descriptor, respectively. the descriptor similarity is evaluated using the hamming distance that computes how many bits of two given feature vectors are different. 5. evaluation the main motivation behind the utilisation of fpga– based systems in visual navigation tasks is a latency reduction in the robotic navigation system which is especially recognizable in experiments performed in dynamic or confined environments, which impose imminent threats to the mobile robotic platform [1, 23] and the reduction of the computational load thus power consumption of the mobile robot platforms. this section presents the results of the experimental evaluation of the impact of the fpga–based coprocessor utilization on the reduction of the computational burden of the feature extraction process in the monocular vision–based teach–and–repeat autonomous navigation. the particular parts of the whole navigation system under the evaluation are the feature extraction chain and the computationally most demanding parts of the navigation algorithm: the feature detection and description; features matching to the map previously created in the teach mode; construction of the navigational histogram; and determination of the robot heading correction based on the maximum peak in the histogram voting method. for the feature extraction the fast feature detector [6] is used with the setting of n = 12, while 256 bit long brief [12] (brief32) is utilized as the feature descriptor. the number of features is set to approx. 200 features per image and the pre-learned map in the navigation pipeline evaluation contains 200 features. the pure cpu based implementation running on the odroid u3 board [24] was compared to the (a) . outdoor urban env. (b) . indoor lab env. figure 4. testing environments. fpga–based soc implementation on terasic de0nano board [25]. in particular, the hardware and software used in the evaluation are as follows: odroid u3 – a small embedded micro-computer suitable for robotic applications. it features 1.7 ghz arm cortex a9 quad-core microprocessor (samsung exynos4412 prime) running the ubuntu 12.04 linux operating system. the attached camera is the mobius actioncam with the resolution of 640×480 with 60 fps. the implementation of the navigation algorithm was based on the open computer vision library [26] (opencv 2.4.9) implementations of the fast feature detector and brief32 descriptor. de0-nano board – an embedded module with the altera cyclone iv ep4ce22 fpga equipped with the aptina mt9v034 grayscale camera sensor with the global shutter and the resolution of 752×480 providing images with 60 fps. the implementation of the navigation algorithm follows the approach described in [20]. it utilizes the bare-metal programmed nios iie softcore processor clocked at 150 mhz for the feature description, matching, and histogram voting and a dedicated fpga–based parallel co-processor for the fast feature detection. the proposed evaluation consists of the focused examination of the feature detector and descriptor and examination of the performance of the whole navigation system. the tests were performed in field conditions both in outdoor urban (fig. 4a) and indoor lab (fig. 4b) environments. first, we evaluated the performance of the fpga– based and cpu–based implementations of the fast feature detector and computation of the brief descriptor. two criteria are considered in this evaluation: correctness of the fpga–based implementation and computational requirements. regarding the correctness of the fpga–based implementation, it provides identical results to the cpu– based counterpart for the detected features and also the same descriptors. this test was performed on artificially generated data in order to ensure the correctness of the results. the data consists of a checkerboard pattern with grayscale gradient which allows to detect various number of corners depending on the predefined threshold. concerning the computational requirements of the 12 vol. 2/2015 on fpga based acceleration of image processing in mobile robotics feature extractor, the odroid implementation needs (in average) 17.07 ms for approximately 200 fast feature detections with brief32 descriptors in an image with the resolution of 640×480 pixels. the de0nano implementation exhibits similar performance with the processing time of 16.91 ms for 200 feature detections and descriptions. the navigation stack performance benchmarking consists of the feature extraction, feature matching against the pre-learned map, and the histogram voting method to establish the heading correction of the robot. the achieved results are summarized in table 1. platform odroid u3 de0-nano cpu cores 4 1 cpu clock 1.7 ghz 150 mhz fpga 22320∗∗ fpga usage 28% feature extraction 17.07 ms 16.91 ms navigation pipeline 35.89 ms 24.38 ms power consumption 8.13 w 1.82 w ∗∗ the number of available logic elements. table 1. navigation pipeline processing results of odroid u3 and de0-nano soc platforms 5.1. discussion the results presented in table 1 indicate that the fpga–based soc approach does not significantly reduce the required time for feature extraction; however, the power consumption is reduced by 4.5 times magnitude in comparison to the purely cpu–based implementation. regarding the update rate of individual methods the fpga–based implementation is capable to process each other image in the 60 fps camera stream, which gives about 30 processed images per second while the cpu–based implementation can process 20 frames per second if we consider a uniform image retrieval; otherwise the update rate of the cpu–based implementation is 27 fps but images are dropped at non-deterministic way which leads to uneven distribution of information in time. the fast feature detector of the de0-nano implementation performs feature detection online on the stream of camera data, which implies that the image coordinates of all salient points in the image are known during the readout of a single image from the camera sensor. the feature description and matching to the pre-learned map is performed in an instant when the feature is detected in the camera stream. however, 256 bit long brief32 descriptor calculation together with matching to approx. 200 features from pre-learned map on the 150 mhz single core cpu makes it impossible to finish all the computations during the readout of one single image. therefore, each other frame in a 60 hz camera stream is skipped, which allows the algorithm to finish all the calculations in less than 32 ms during the readout of the next image. 6. conclusion in this paper, we consider a problem of decreasing the computational load of the embedded computational resources utilized in the vision-based navigation tasks. the presented results show that although the processing times of both the fpga–based and cpu– based implementations are almost similar most-likely due to the overall low complexity of the used feature extraction, the fpga–based system–on–chip design can significantly reduce the power consumption of the embedded processor in comparison to a purely cpu–based solution. in order to further verify and quantify the results, we plan to thoroughly benchmark each part of the visual navigation algorithm stack and incorporate the fpga-based image processing in a more complex navigation task of a visual slam, where the computational power of the fpga platform can be of potential benefit. acknowledgements the presented work has been supported by the czech science foundation (gačr) under research project no. 13-18316p. references [1] n. michael, d. mellinger, q. lindsey, v. kumar. the grasp multiple micro-uav testbed. robotics automation magazine, ieee 17(3):56–65, 2010. doi:10.1109/mra.2010.937855. 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[20] p. čížek. embedded module for image processing. master’s thesis, ctu, faculty of electrical engineering, 2015. [21] garcía, gabriel j and jara, carlos a and pomares, jorge and alabdo, aiman and poggi, lucas m and torres, fernando. a survey on fpga-based sensor systems: towards intelligent and reconfigurable low-power sensors for computer vision, control and signal processing. sensors 14(4):6247–6278, 2014. doi:10.3390/s140406247. [22] t. krajník, p. decristoforis, m. nitsche, et al. image features and seasons revisited. ecmr 2015 – to appear . [23] s. lupashin, m. hehn, m. w. mueller, et al. a platform for aerial robotics research and demonstration: the flying machine arena. mechatronics 24(1):41–54, 2014. doi:10.1016/j.mechatronics.2013.11.006. [24] collective of authors. hardkernel co., ltd.@online. http://www.hardkernel.com, cited on 2015/08/17. [25] collective of authors. terasic, inc.@online. http://www.terasic.com.tw, cited on 2015/08/17. [26] bradski, g. and kaebler, a. computer vision with the opencv library. o’reilly media, 2008. 14 http://dx.doi.org/10.1109/ismar.2007.4538852 http://dx.doi.org/10.1109/iccv.2011.6126542 http://dx.doi.org/10.1109/iccv.2011.6126544 http://dx.doi.org/10.1002/rob.20354 http://dx.doi.org/10.1007/978-3-642-15561-1_56 http://dx.doi.org/10.1007/s00138-014-0599-0 http://dx.doi.org/10.1016/j.cviu.2007.09.014 http://dx.doi.org/10.1109/iros.2014.6943263 http://dx.doi.org/10.1109/icra.2015.7138979 http://dx.doi.org/10.1109/icra.2014.6907625 http://dx.doi.org/10.1109/icra.2014.6906892 http://dx.doi.org/10.1109/icra.2014.6907232 http://dx.doi.org/10.3390/s140406247 http://dx.doi.org/10.1016/j.mechatronics.2013.11.006 http://www.hardkernel.com http://www.terasic.com.tw acta polytechnica ctu proceedings 2:8–14, 2015 1 introduction 2 related work 3 processor centric design 4 bearing-only visual navigation 4.1 features from accelerated segment test 4.2 binary robust independent elementary features 5 evaluation 5.1 discussion 6 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0001 acta polytechnica ctu proceedings 4:1–7, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app study of identification of two-phase flow parameters by pressure fluctuation analysis ondrej burian∗, vaclav dostal, ladislav vesely faculty of mechanical engineering, czech technical university in prague, technická 4, praha 6, czech republic ∗ corresponding author: ondrej.burian@fs.cvut.cz abstract. this paper deals with an identification of a parameters of simple a pool boiling in a vertical rectangular channel by analysis of pressure fluctuation. a small experimental facility about 9 kw power, which was used for simulation of apool boiling phenomena and creation of steam-water volume in this work is introduced. a several pressure fluctuations measurements and differential pressure fluctuations measurements at various were carried out. main changed parameters were a power of heaters and a hydraulics resistance of a channel internals. a measured pressure data were statistically analysed and compared with goal to find dependencies between parameters of a two-phase flow and statistical properties of pressure fluctuation. at the end of this paper are summarized final results and applicability of this method for parameters determination of two phase flow for pool boiling conditions at ambient pressure is discused. keywords: two-phase flow, pressure fluctuation measurement, void fraction. 1. introduction boiling of liquids is a high dynamic phenomena, which is characterized as a phase change of a liquid phase to a gaseous phase. the most common example of that is boiling of water to water steam, which occurs in many facilities in our life. it may exist in wide range of parameters and regiments, which includes pressure, temperature, heat flux, bulk temperature etc. the one of the key parameters is a void fraction which represents a share of steam fraction in overall volume. the knowledge of this parameter is key for identification other parameters of boiling and operational parameters of facilities which operate with boiling phenomena like steam boilers, boiling water reactors or steam generators. one of a perspective method for identification a key parameters of boiling and two-phase flow phenomena, which is focused on analysis of measured differential pressure and its a time fluctuation, is described in this work. this method is based on fact that differential pressure between two measured locations in a two-phase mixture located in some distance at above is dependent on the void fraction. void fraction in volumetric form is defined as a ratio of a steam fraction volume to overall a two-phase volume, see (1). the density of a two-phase mixture can be expressed by (2), where ρl is density of liquid fraction and ρg is density of steam fraction. α = vg v (1) ρm = ρl(1 −〈α〉) + ρg〈α〉 (2) second assumption is that the value of this differential pressure is not stable in stable state, but it fluctuated around some mean value. this dependency is a result of influence of the void fraction on a hydrostatic pressure at measured locations – a measurement points. hydrostatic pressure of a two-phase volume about some void fraction is value between hydrostatic pressure of a liquid volume and a steam volume, which decreases with increasing a void fraction. a pressure fluctuation is a result of a process of a bubble generation and its flow throw the measured section. if we find dependency between void fraction and measured differential pressure, which is measured between two measurement points located at above in some distance, we can use it like fundamental of a method for determination of the void fraction. however more interesting may be an information about boiling and two-phase flow parameters which we can get from statistical analysis of fluctuation of measured differential pressure. research of this method, identification advantages and disadvantages of this method is subject of this research work which will be introduced in this paper. due to a boiling and a two-phase flow are complicated phenomenas, is impossible to determinate an analytical solution with physical model of dependency between void fraction and differential pressure. they are complex phenomenas which are affected many external and internal influences. from that reasons as described at above was chose experimental way which should give empirical view of scope to this problem. using differential pressure as a parameter is more useful than point pressure, because we do not know position of a level of a two-phase mixture up the measured points. the beginning phase of this research project is focused on an experimental research at an ambient pressure conditions. the aim of this phase is a verification an initial assumptions, assessment a general values of 1 http://dx.doi.org/10.14311/ap.2016.4.0001 http://ojs.cvut.cz/ojs/index.php/app 230 acta polytechnica ctu proceedings 2(1): 230–233, 2015 230 doi: 10.14311/app.2015.02.0230 life after eruption: best of 2009–2013 c. tappert1, a. ederoclite2, l. schmidtobreick3, n. vogt1 1departamento de f́ısica y astronomı́a, universidad de valparáıso, chile 2centro de estudios de f́ısica del cosmos de aragón, teruel, spain 3european southern observatory, santiago, chile corresponding author: claus.tappert@uv.cl abstract from our ongoing survey to study the post-nova population we present details on the four objects v728 sco, ar cir, v972 oph and x cir. keywords: cataclysmic variables novae individual: v728 sco ar cir v972 oph x cir. 1 introduction in 2009 we have begun a survey on post-novae that had erupted before 1980, in order to address the lack of identified post-nova systems. scope, methods and the current state of the survey are presented by ederoclite et al. elsewhere in this volume. in the present article we take a closer look at four objects from our survey with peculiar properties that deserve further investigations. a detailed analysis of three of these systems, v728 sco, ar cir and v972 oph, can be found in tappert et al. (2013a, b), and we here only briefly summarize the results. 2 v728 sco the cv corresponding to nova sco 1862 was recovered as a v = 18.5 mag object roughly 2 arcmin north-west of its suspected position (tappert et al. 2012). its spectrum presents – for an old nova – unusually strong emission lines. photometric time-series revealed the object to be eclipsing with an orbital period porb = 3.32 h. this places v728 sco within the regime of the sw sex type systems, cvs that are characterized by very high mass-transfer rates ṁ (e.g., rodŕıguez-gil et al. 2007a, b; see also schmidtobreick & tappert, this volume). furthermore, old novae in general are expected to run high ṁ due to irradiation of the secondary star by the eruption-heated white dwarf (kovetz et al. 1988), and in general this is confirmed by observations (iben et al. 1992). for example, the old nova rr pic, with porb = 3.48 h, does indeed share the characteristics of an sw sex star (schmidtobreick et al. 2003a, 2008). v728 sco, on the other hand, appears to have a lower ṁ than expected. evidence for this does not only stem from the spectroscopic appearance, but also from the long-term behaviour that shows the system to vary between a low and a ∼1.5 mag brighter state. while the data coverage is still very sparse, it looks like this brighter state is in general short-lived and bears resemblance to the outbursts of dwarf novae. analysis of one eclipse that was observed in the low state provided an estimate of the radius of the eclipsed source to r = 0.09 r�, which is significantly larger than any feasible solution for a white dwarf. we interpret this result as the presence of a hot inner disc, such as has been invoked by schreiber et al. (2000) to explain the low-amplitude high-frequency outbursts that are observed in some post-novae and nova-like stars (e.g., honeycutt et al. 1998), and would also fit the longterm behaviour of v728 sco. our eclipse data provide the first direct evidence for the presence of such disc. 3 ar cir ar cir was a comparatively slow nova that erupted in 1906. it was recovered by duerbeck & grebel (1993). time-series spectroscopy taken by us on three consecutive nights in 2012 yielded porb = 5.14 h. the spectrum shows a moderately strong hα emission line on a red continuum, the latter being very probably entirely caused by interstellar reddening rather than due to the contribution of the secondary star. the r-band brightness shows a clear modulation with porb and an amplitude of ∼0.3 mag, indicating that the inclination is comparatively high. the hα line profile shows an intriguing variability in that at certain orbital phases, and strongest at phase 0.5 (corresponding to the maximum of the r-band light curve), a p cyg component becomes visible in the form of a high-velocity blue absorption and a red emission component. we interpret 230 http://dx.doi.org/10.14311/app.2015.02.0230 life after eruption: best of 2009–2013 this as an optically thick outflow and determine a rough estimate of the projected outflow velocity to 700 ± 100 km/s. 4 v972 oph time-series spectroscopic observations of this post-nova (nova oph 1957) yield an orbital period porb = 6.75 h thus making it a member of the minority of novae with comparatively long orbital periods (see ederoclite et al., this volume). examination of the profile of the hα emission line reveals an additional emission at ∼6580 å. this component moves roughly parallel to the main hα component. this practically excludes scenarios of an additional hα component from the secondary star or from the hot spot, as well as a nebular line like [nii] λ6584. the absence of an absorption line blueward of hα lets an origin in an outflow, such as assumed for ar cir, also appear unlikely. instead we tentatively identify that line with the cii λλ6578/6583 doublet. to our knowledge, this is the first time that this line is observed in a cv in this form (it had also be identified as originating in the outflowing material of the helium nova v445 pup; iijima & nakanashi 2008). while carbon emission has been observed before in some post-novae (e.g., schmidtobreick et al. 2003b, bianchini et al. 2012), this concerns mostly the blue part of the spectrum. unfortunately the available spectra of v972 oph by zwitter & munari (1996) and ringwald et al. (1996) that cover that range have too low s/n to confirm the presence of such lines in v972 oph. the weakness of the hα emission line (which has an equivalent width of 2 å) certainly helped in the detection of the cii line. this raises the question whether this line is not more frequently present in cvs, but hidden away in the wings of a stronger hα emission. a quick examination of seven novae finds three more systems with likely excess emission at ∼6580 å. as a caveat we note that this list includes ar cir, where we find an outflow as a more likely explanation. however, the other systems lack an accompanying blue emission or absorption component that would suggest such phenomenon as the origin. 5 x cir the eruption of nova cir 1927 was reported by becker (1929). the aftermath of the nova is little studied. duerbeck (1987) classifies it as a slow nova with t3 = 170 d. woudt & warner (2002) identified a possible candidate for the post-nova due to short-term variability, but a spectrum taken by mason & howell (2003) revealed that object to have a reddish continuum without emission lines. our ubvr photometry from may 2009 (fig. 1, top) allows for a number of candidates based on colour and distance from the reported coordinates. in 2012 we obtained spectroscopic data for five of them with the mos mode of fors2 at the eso-vlt (ut1). the post-nova was confirmed to be an object roughly 1.2 arcmin north-west of its suspected position. the spectrum shows comparatively strong emission lines (equivalent width of hα ∼30 å) of the balmer and hei series (fig. 1, bottom). 0.0 0.5 1.0 1.5 v −r [mag] −0.5 0.0 0.5 1.0 1.5 u − b [m ag ] 4000 5000 6000 7000 8000 wavelength [å] 5 10 f lu x [1 0 − 1 7 er g cm − 1 s− 1 å − 1 ] figure 1: top: colour-colour diagram of the field (4.5×4.5 arcmin) of x cir. the confirmed post-nova is marked by a circle. bottom: low-resolution spectrum of the post-nova. while being situated in a promising place in the colour-colour diagram, the later confirmed post-nova was not our best candidate, but in fact rather represented the very bottom of our list. the reason for this is that our photometric data showed a larger point-spread function (psf) than for the other stars in the field, indicating an extended object. this was later confirmed by the 2d spectroscopic data. an examination of the latter showed that the emission lines appear to be confined to the one part of the spectroscopic psf that corresponds to the eastern part of the photometric psf. our preliminary conclusion is therefore that the extension is not due to a shell, but that this is an unresolved (< 0.5 arcsec) visual binary. the spectrum of x cir presents a prominent heii 4686 line, and even, albeit much weaker, heii 5411 (fig. 2). such strong presence of heii is often a signature of a magnetic system. on the other hand, the hei lines are distinctively double-peaked with a deep central valley, indicating the presence of an accretion disc. a possible explanation is that x cir is an intermediate polar. the heii 5411 line is also a transient feature. in our three spectra, taken with a time separation of ∼3 h between the first and the second, and ∼24 h between the second and the third, the line is strongest in the first (which is the one in fig. 2), not detected in 231 c. tappert et al. the second, and weakly present in the third spectrum (fig. 1). since the individual integration times amount to 15 min, this could be an orbital effect. the heii lines and the balmer lines are single-peaked, suggesting that at least a major part of their profile has an origin different from that of hei. 5400 5600 5800 wavelength [å] 1.0 1.2 n or m al is ed f lu x heii hei? ? 7000 7200 7400 wavelength [å] hei ? figure 2: close-up on the spectrum of x cir. the label mark the line identifications; a question mark is used for the three unidentified lines. note the deep central valleys in the double-peaked profile of the two present hei lines. we furthermore find three emission lines that we have not yet been able to identify (fig. 2). taking into account the observed shift of the identified lines, we estimate their corresponding rest wavelengths to ≥5678, ≥5805, ≥7233, respectively. examining lists of spectral lines (e.g., coluzzi 1999) does not present obvious suspects for an identification. we note that the profiles are also single-peaked, so that these probably do not originate exclusively in the accretion disc. we derive the ’quiescent’ brightness of x cir in the v filter to 19.3 mag. the reported photographic value of the eruption maximum is 6.5 mag (duerbeck 1987). assuming that the effect of the difference between the filters is < 0.5 mag, this yields an eruption amplitude > 12 mag. this places it a bit off the fit to the rateof-decline / eruption-amplitude relation (ederoclite et al., this volume; note that x cir is not yet included in their figure), but still within the larger scatter. 6 summary we have presented four post-novae with apparently peculiar properties. v728 sco is one of the few objects that defies the general trend that post-novae are cvs with very high mass-transfer rates (ṁ). finding such objects will be very important to answer the question whether post-novae in general have high ṁ as a consequence of the nova eruption, or if high ṁ cvs are just more likely to erupt as a nova. v728 sco is furthermore deeply eclipsing, which presents an excellent and rarely found opportunity to determine accurate system parameters of a nova. ar cir, on the other hand, while with more than 100 yr having passed since its eruption can already be counted among the ’old’ old novae, apparently still drives such high ṁ that it causes an outflow of material that is visible in the optical spectrum. since our data only covers the spectral range next to hα it would be interesting to examine other spectral ranges and lines for this phenomenon to gain more information on the outflow. in v972 oph we detect an unusual emission line that we attribute to carbon. we argue that such line could in principle be more frequent in cvs because it can be easily hidden in the wings of the usually strong hα emission that is exceptionally weak in v972 oph. the currently available data on the blue spectral range has too low s/n to examine it for other tracers of the presence of carbon in v972 oph, thus a higher quality spectrum of that region would be most welcome. the carbon line in v972 oph moves parallel to hα and thus likely originates in the accretion disk or the bright spot, or on the secondary star itself. in any case it represents material from the secondary. it was either accreted from the expelled material during the nova eruption or produced in the secondary, which then could not have formed as a late-type main-sequence star, but must have undergone some nuclear evolution prior to its semi-detached cv phase. thus, the study of carbon in (not only) post-novae bears relevance for our understanding on cv evolution. the results on x cir are still somewhat preliminary. the strong balmer lines argue for a comparatively low ṁ, the prominent presence of heii could point to a magnetic white dwarf, but at least is evidence for a hot region somewhere in the system, while the double-peaked hei profiles indicate the presence of an accretion disc. a number of emission features yet remain to be identified which means that these lines are at least very rarely observed in cvs. finally, the spectral appearance, the transient nature of heii 5411 and the double-peaked hei lines suggest a high inclination. there is therefore a good chance that the orbital period will be accessible comparatively easily via time-series photometry. we hope that this presentation motivates further studies of those systems. while their properties may appear peculiar, they might not be exclusive to those objects. because the number of detailed studies on post-novae is still fairly small, we cannot yet be certain if they are more representative of the rule or the exception. acknowledgement this research was supported by fondecyt regular grant 1120338 (ct and nv). ae acknowledges support by the spanish plan nacional de astrononomı́a y astrof́ısica under grant aya2011-29517-c03-01. 232 life after eruption: best of 2009–2013 references [1] becker, f.: 1929, astron. nachr., 237, 71. [2] bianchini, a., saygac, t., orio, m., et al.: 2012, a&a, 539, a94. [3] coluzzi, r.: 1999, vizier online data catalog, 6071. [4] duerbeck, h.w.: 1987, space sci. rev., 45, 1. doi:10.1007/bf00187826 [5] duerbeck, h.w., grebel, e.k.: 1993, mnras, 265, l9. doi:10.1093/mnras/265.1.l9 [6] honeycutt, r.k., robertson, j.w., turner, g.w.: 1998, aj, 115, 2527. [7] iben, jr., i., fujimoto, m.y., macdonald, j.: 1992, apj, 384, 580. doi:10.1086/170900 [8] iijima, t., nakanashi, h.: 2008, a&a, 482, 865. [9] kovetz, a., prialnik, d., shara, m.m.: 1988, apj, 325, 828. [10] mason, e., howell, s.b.: 2003, a&a, 403, 699. [11] ringwald, f.a., naylor, t., mukai, k.: 1996, mnras, 281, 192. [12] rodŕıguez-gil, p., schmidtobreick, l., gänsicke, b.t.: 2007a, mnras, 374, 1359. doi:10.1111/j.1365-2966.2006.11245.x [13] rodŕıguez-gil, p., gänsicke, b.t., hagen, h.-j., et al.: 2007b, mnras, 377, 1747. doi:10.1111/j.1365-2966.2007.11743.x [14] schmidtobreick, l., tappert, c., saviane, i.: 2003a, mnras, 342, 145. doi:10.1046/j.1365-8711.2003.06523.x [15] schmidtobreick, l., tappert, c., bianchini, a., mennickent, r.e.: 2003b, a&a, 410, 943. [16] schmidtobreick, l., papadaki, c., tappert, c., ederoclite, a.: 2008, mnras, 389, 1345. doi:10.1111/j.1365-2966.2008.13641.x [17] schreiber, m.r. and gänsicke, b.t. and cannizzo, j.k.: 2000, a&a, 362, 268. [18] tappert, c., ederoclite, a., mennickent, r.e., et al.: 2012, mnras, 423, 2476. [19] tappert, c., vogt, n., schmidtobreick, l., et al.: 2013a, mnras, 431, 92. [20] tappert, c., schmidtobreick, l., vogt. n., ederoclite, a.: 2013b, mnras, in press (arxiv 1302.5570). [21] woudt, p.a., warner, b.: 2002, in: classical nova explosions, hernanz, m. & josé, j., eds., aip conf. proc., vol. 637, p.532. [22] zwitter, t., munari, u.: 1996, a&as, 117, 449. discussion christian knigge: a comment on v728 sco: i think this is totally consistent with being a ”normal” sw sex star. it has the same v-shaped eclipse in the high state (and may be self-occulting), and several sw sex stars, like dw uma, also show high-state / lowstate behaviour. claus tappert: it is true that the high-state eclipse very much resembles that of sw sex stars. but i would argue that the other properties do not. we do not see any high-velocity wings (this might arguably be due to the quality of the spectra), and the emission lines are distinctively double-peaked. we need more long-term data to see if the system spends more time in low or high state, but so far my best bet would be that v728 sco is in a somewhat transitory phase, not quite sw sex, not quite dwarf nova. 233 http://dx.doi.org/10.1007/bf00187826 http://dx.doi.org/10.1093/mnras/265.1.l9 http://dx.doi.org/10.1086/170900 http://dx.doi.org/10.1111/j.1365-2966.2006.11245.x http://dx.doi.org/10.1111/j.1365-2966.2007.11743.x http://dx.doi.org/10.1046/j.1365-8711.2003.06523.x http://dx.doi.org/10.1111/j.1365-2966.2008.13641.x introduction v728 sco ar cir v972 oph x cir summary 71 acta polytechnica ctu proceedings 1(1): 71–78, 2014 71 doi: 10.14311/app.2014.01.0071 clustering measurements of broad-line agns: review and future mirko krumpe1,3, takamitsu miyaji2,3, and alison l. coil3 1european southern observatory, karl-schwarzschild-straße 2, 85748 garching bei münchen, germany 2instituto de astronomı́a, unam, apdo. postal 106, ensenada, bc, méxico 3university of california san diego, cass, 9500 gilman drive, la jolla, ca 92093-0424, usa corresponding author: mkrumpe@eso.org abstract despite substantial effort, the precise physical processes that lead to the growth of super-massive black holes in the centers of galaxies are still not well understood. these phases of black hole growth are thought to be of key importance in understanding galaxy evolution. forthcoming missions such as erosita, hetdex, eboss, bigboss, lsst, and panstarrs will compile by far the largest ever active galactic nuclei (agns) catalogs which will allow us to measure the spatial distribution of agns in the universe with unprecedented accuracy. for the first time, agn clustering measurements will reach a level of precision that will not only allow for an alternative approach to answering open questions in agn and galaxy co-evolution but will open a new frontier, allowing us to precisely determine cosmological parameters. this paper reviews large-scale clustering measurements of broad line agns. we summarize how clustering is measured and which constraints can be derived from agn clustering measurements, we discuss recent developments, and we briefly describe future projects that will deliver extremely large agn samples which will enable agn clustering measurements of unprecedented accuracy. in order to maximize the scientific return on the research fields of agn and galaxy evolution and cosmology, we advise that the community develops a full understanding of the systematic uncertainties which will, in contrast to today’s measurement, be the dominant source of uncertainty. keywords: large-scale structure of universe galaxies: active. 1 introduction large area surveys such as the two degree field galaxy redshift survey (2dfgrs; colless et al. 2001) and the sloan digital sky survey (sdss; abazajian et al. 2009) have measured positions and redshifts of millions of galaxies. these measurements allow us to map the 3d structure of the nearby universe1. galaxies are not randomly distributed in space. they form a complex cosmic network of galaxy clusters, groups, filaments, isolated field galaxies, and voids, which are large regions of space that are almost devoid of galaxies. the current understanding of the distribution of galaxies and structure formation in the universe is based on the theory of gravitational instability. very early density fluctuations became the “seeds” of cosmic structure. these have been observed as small temperature fluctuations (δt/t ∼ 5 × 10−5) in the cosmic microwave background with the cosmic background explorer (smoot et al. 1992). the small primordial matter density enhancements have progressively grown through gravitational collapse and created the complex network seen in the distribution of matter in the later universe. during a galaxy’s lifetime different physical processes, which are still not well understood, can trigger a mass flow onto the central super-massive black hole (smbh). in this phase of galaxy evolution, the galaxy is observed as an active galactic nucleus (agn). after several million years, when the smbh has consumed its accretion reservoir, the central engine shuts down, and the object is again observed as a normal galaxy. the agn phase is thought to be a repeating special epoch in the process of galaxy evolution. in recent years it has become evident that both fundamental galaxy and agn parameters change significantly between low (z < 0.3) and intermediate redshifts (z ∼ 1 − 2), e.g., global star formation density (hopkins & beacom 2006) and accretion rate onto smbhs. for example, the contribution to black hole growth has shifted from high luminosity objects at high redshifts to low luminosity objects at low redshifts (agn “downsizing”; e.g., hasinger et al. 2005). it has also become clear that smbh masses follow a tight relation with the mass or velocity dispersion of the stars in galactic bulges (magorrian et al. 1998; gebhardt et al. 2000; ferrarese & merritt 2000). these observational correlations moti1a visual impression is given in this video: http://vimeo.com/4169279 71 http://dx.doi.org/10.14311/app.2014.01.0071 mirko krumpe, takamitsu miyaji, alison l. coil vate a co-evolution scenario for galaxies and agns and provide evidence of a possible interaction or feedback mechanism between the smbh and the host galaxy. the interpretation of this correlation, i.e., whether and to what extent the agn influences its host galaxy, remains controversial (e.g., jahnke & macció 2011). since agns are generally much brighter than (inactive) galaxies, one major advantage of agn large-scale (i.e., larger than the size of a galaxy) clustering measurements over galaxy clustering measurements is that they allow the study of the matter distribution in the universe out to higher redshifts. at these redshifts, it becomes challenging and observationally expensive to detect galaxies in sufficient numbers. furthermore, as the distribution of agns and galaxies in the universe depends on galaxy evolution physics, large-scale clustering measurements are an independent method to identify and constrain the physical processes that turn an inactive galaxy into an agn and are responsible for agn and galaxy co-evolution. in the last decade the scientific interest in agn large-scale clustering measurements has increased significantly. as only a very small fraction of galaxies contain an agn (∼1%), the remaining and dominating challenge in deriving physical constraints based on agn clustering measurements is the relative small sample size compared to galaxy clustering measurements. however, this situation will change entirely in the next decade when several different surveys come online that are expected to identify millions of agn over ∼80% of cosmic time. we therefore review broad-line agn clustering measurements. a general introduction to clustering measurements is given in sections 2 & 3. in section 4 we briefly summarize how agn clustering measurements have evolved and discuss recent developments. in section 5 we discuss the outlook for agn clustering measurements in future upcoming projects. 2 understanding observed clustering properties in our current understanding, the observed galaxy and agn spatial distribution in the universe – i.e., largescale clustering – is caused by the interplay between cosmology and the physics of galaxy evolution. in the commonly assumed standard cosmological model, lambda-cdm, the universe is currently composed of ∼70% dark energy, ∼25% dark matter (dm), and ∼5% baryonic matter (larsen et al. 2011). dark matter plays a key role in structure formation as it is the dominant form of matter in the universe. baryonic matter settles in the deep gravitational potentials created by dark matter, the so-called dark matter halos (dmhs). the term “halo” commonly refers to a bound, gravitationally collapsed dark matter structure which is approximately in dynamical equilibrium. the parameters of the cosmological model determine how the dmhs are distributed in space (fig. 1, left panel, a-branch) as a function of the dmh mass and cosmic time. different cosmological models lead to different properties of the dmh population. figure 1: current conceptual model of the physical processes involved in large-scale galaxy and agn clustering. left: the two branches (a and b) in the diagram show the primary causes of clustering: (a) the properties of the dark matter halo population, which are based on the cosmological model, and (b) the physics of complex processes in galaxy formation and evolution, which lead to a distinct baryonic population within collapsed dark matter halos. figure adapted from weinberg 2002. right: illustration of the spatial distribution of galaxies within a dark matter halo. the picture maps the galaxy cluster abell 1689, where an optical image showing the galaxy cluster members is superimposed with the distribution of dark matter shown in purple. credit: nasa, esa, e. jullo, p. natarajan, and j-p. kneib. inside dmhs, or within halos inside another dmh, called sub-halos, the baryonic gas will radiatively cool. if the gas reservoir is large enough, star and galaxy formation will be initiated. the gas can also be accreted onto the smbh in the center of the galaxy. on scales comparable to the size of the galaxy, the agn can heat and/or eject the surrounding gas, preventing star formation, and eventually removing the gas fueling the agn itself. all the galaxy evolution processes described here determine how galaxies and agns are distributed within dmhs (fig. 1, left panel, b-branch). this distribution of agn and galaxies within dmhs (fig. 1, right panel) is described by the halo occupation distribution (hod; peacock & smith 2000). in addition to the spatial distribution of agn and galaxies in dmhs, the hod describes the probability distributions 72 clustering measurements of broad-line agns: review and future of the number of agns and galaxies per dmh of a certain mass and the velocity distribution of agns and galaxies within a dmh. the interplay between cosmology and galaxy evolution causes the observed large-scale clustering of galaxies and agns. the goal of agn and galaxy clustering measurements is to reverse the causal arrows in the fig. 1 (left panel), working backwards from the data to the galaxy & agn halo occupation distribution and dmh population properties, in order to finally draw conclusions about galaxy and agn physics, as well as to constrain fundamental cosmological parameters. 3 clustering measurements the most common statistical estimator for large-scale clustering is the two-point correlation function (2pcf; peebles 1980) ξ(r). this quantity measures the spatial clustering of a class of object in excess of a poisson distribution. in practice, ξ(r) is obtained by counting pairs of objects with a given separation and comparing them to the number of pairs in a random sample with the same separation. different correlation estimators are described in the literature (e.g., davis & peebles 1983; landy & szalay 1993). the large-scale clustering of a given class of objects can be quantified by computing the angular (2d) correlation function, which is the projection onto the plane of the sky, or with the spatial (3d) correlation function, which requires redshift information for each object. obtaining spectra to measure the 3d correlation function is observationally expensive, which is the main reason why some studies have had to rely on angular correlation functions. however, 3d correlation function measurements are by far preferable, since the deprojection (limber 1954) of the angular correlation function introduces large systematic uncertainties. despite these large caveats and the already moderately low uncertainties of current 3d correlation measurements, the use of angular correlation functions might still be justified when exploring a new parameter space. however, the next generation multi-object spectrographs (e.g., 4most (de jong et al. 2012), bigboss (schlegel et al. 2011), and weave (dalton et al. 2012)), will make it far easier to simultaneously obtain thousands of spectra over wide fields. hence, measurements of the 3d correlation function will soon become ubiquitous. as one measures line-of-sight distances for 3d correlation functions from redshifts, measurements of ξ(r) are affected by redshift-space distortions due to peculiar velocities of the objects within dmhs. to remove this effect, ξ(r) is commonly extracted by counting pairs on a 2d grid of separations where rp is perpendicular to the line of sight and π is along the line of sight. then, integrating along the π-direction leads to the projected correlation function, wp(rp), which is free of redshift distortions. the 3d correlation function ξ(r) can be recovered from the projected correlation function (davis & peebles 1983). the resulting signal can be approximated by a power law where the largest clustering strength is found at small scales. at large separations of >50 mpc h−1 the distribution of objects in the universe becomes nearly indistinguishable from a randomly-distributed sample. only on comoving scales of ∼100 mpc h−1 can a weak positive signal be detected (e.g., eisenstein et al. 2005; cole et al. 2005) which is caused by baryonic acoustic oscillations (bao) in the early universe. the spatial clustering of observable objects does not precisely mirror the clustering of matter in the universe. in general, the large-scale density distribution of an object class is a function of the underlying dark matter density. this relation of how an object class traces the underlying dark matter density is quantified using the linear bias parameter b. this contrast enhancement factor is the ratio of the mean overdensity of the observable object class, the so-called tracer set, to the mean overdensity of the dark matter field, defined as b = (δρ/〈ρ〉)tracer/(δρ/〈ρ〉)dm, where δρ = ρ−〈ρ〉, ρ is the local mass density, and 〈ρ〉 is the mean mass density on that scale. in terms of the correlation function, the bias parameter is defined as the square root of the 2pcf ratio of the tracer set to the dark matter field: b = √ ξtracer/ξdm. rare objects which form only in the highest density peaks of the mass distribution have a large bias parameter and consequently a large clustering strength. theoretical studies of dmhs (e.g., mo & white 1996; sheth et al. 2001) have established a solid understanding of the bias parameter of dmhs with respect to various parameters. comparing the bias parameter of an object class with that of dmhs in a certain mass range at the same cosmological epoch allows one to determine the dmh mass which hosts the object class of interest. a halo may contain substructures, but the dmh mass inferred from the linear bias parameter refers to the single, largest (parent) halo in the context of hod models. 3.1 why are we interested in agn clustering? agn clustering measurements explore different physics on different scales. at scales up to the typical size of a dmh (∼ 1 − 2 mpc), clustering measurements are sensitive to the physics of galaxy and agn formation and evolution. constraints on the galaxy and agn merger rate and the radial distribution of these objects within dmhs can be derived. on scales larger than the size of dmhs, the large-scale clustering is sensi73 mirko krumpe, takamitsu miyaji, alison l. coil tive to the underlying dm density field, which essentially depends only on cosmological parameters. consequently, with only one measurement both galaxy and agn co-evolution as well as cosmological parameters can be studied. future high precision agn clustering measurements have the potential to accurately establish missing fundamental parameters that describe when agn activity and feedback occur as a function of luminosity and redshift. since they will precisely determine how dmhs are populated by agn host galaxies, these measurements will also improve our theoretical understanding of galaxy and agn evolution by enabling comparisons to galaxy measurements and cosmological simulations. here, we elaborate on some (though not all) of the critical observational constraints which are provided by agn clustering measurements: • agn host galaxy – agn cannot be more clustered than the type of galaxies they reside in. thus, agn clustering measurements determine the host galaxy type in a statistical sense for the entire agn sample, regardless of the agn’s luminosity. comparing the observed agn clustering to very accurate galaxy clustering measurements, which depend on different galaxy subclasses (morphological, spectral type, luminosity), constrains the agn host galaxy type. • external (mergers) vs. internal triggering – different theoretical models (e.g., fry 1996; sheth et al. 2001; shen 2009) of how agns are triggered predict very different large-scale clustering properties with agn parameters such as luminosity and redshift. moderately precise agn clustering measurements allow us to distinguish between these different models (allevato et al. 2011). furthermore, the validity of different models can be tested for different luminosities and cosmological epochs. • fundamental galaxy/agn physics – agn largescale clustering dependences with various agn properties could potentially be a key source of independent constraints on galaxy/agn physics. comparing the observed agn clustering properties with results from simulations with different inputs for galaxy/agn physics could identify the physics that links the evolution of agns and galaxies. • agn lifetimes – agn clustering measurements allow us to estimate the agn lifetime at different cosmological epochs (martini & weinberg 2001). the underlying idea is that rare, massive dmhs are highly biased tracers of the underlying mass distribution, while more common objects are less strongly biased (kaiser 1984). therefore, if agns are heavily biased they must be in rare, massive dmhs. the ratio of the agn number density to the host halo number density is a measure of the “duty cycle”, i.e., the fraction of the time that the object spends in the agn phase. • cosmological parameters – as agn clustering measurements extend to much higher redshifts than galaxy clustering measurements, they can be used to derive constraints on cosmological parameters (e.g., basilakos & plionis 2009) back to the time of the formation of the first agns. currently, the detection of the bao imprint on clustering measurements at different cosmological epochs is of great interest to constrain the equation of state of dark energy. agn large-scale clustering measurements with very large agn samples can detect the bao signal in a redshift range that is not accessible with galaxy clustering measurements. 4 agn clustering measurements: past and present until the 1980s, studies had to primarily rely on small, optically-selected, very luminous agn samples for clustering measurements. then the main question was whether agns are randomly distributed in the universe (e.g., bolton et al. 1976; setti & woltjer 1977). the extremely small sample sizes did not allow clustering measurements at scales below ∼50 mpc, where a significant deviation from a random distribution is present. thanks to the launch of major x-ray missions in the 1980s and 1990s such as einstein (giacconi et al. 1979) and rosat (truemper 1993), much larger agn samples enabled successful detections of the agn large-scale clustering signal. a detailed review on the history of x-ray agn clustering measurements is given in cappelluti et al. (2012). although agn clustering measurements are far from being as precise as galaxy clustering measurements, some general findings have emerged in recent years. interestingly, over all of studied cosmic time (z ∼ 0 − 3), broad-line agns occupy dmh masses of log (mdmh/[h −1m�]) ∼ 12.0−13.5 and therefore cluster like groups of galaxies. more detailed information about the current picture of broad-line agn clustering is presented in section 6.6 of krumpe et al. (2012). some puzzling questions remain. for example, at z < 0.5 a weak x-ray luminosity dependence on the clustering strength is found (in that luminous x-ray agns cluster more strongly than their low luminosity counterparts, e.g., krumpe et al. 2010; cappelluti et al. 2010; shen et al. 2013). however, at high redshift it seems that high luminosity, optically-selected agns cluster less strongly than moderately-luminous 74 clustering measurements of broad-line agns: review and future x-ray selected agns. whether this finding is due to differences in the agn populations, an intrinsic luminosity dependence to the clustering amplitude, or an observational bias is yet not understood. we note that different studies have used different relations to translate the measured linear bias parameter to dmh mass, as well as different σ8 values. therefore, instead of blindly comparing the derived dmh mass, re-calculating the masses based on the same linear bias to dmh mass relation and the same σ8 is essential when comparing measurements in the literature. 4.1 recent developments in the last few years several new approaches have been used to improve the precision of agn clustering measurements or their interpretation. we summarize these developments below. cross-correlation measurements: auto-correlation function (acf) measurements of broad-line agns often have large uncertainties due to the low number of objects. especially at low redshifts, large galaxy samples with spectroscopic redshifts are frequently available. in such cases, the statistical uncertainties of agn clustering measurements can be reduced significantly by computing the cross-correlation function (ccf). the ccf measures the clustering of objects between two different object classes (e.g., broadline agns and galaxies), while the acf measures the spatial clustering of objects in the same sample (e.g., galaxies or agns). ccfs have been used before to study the dependence of the agn clustering signal with different agn parameters. however, these values could not be compared to other studies as the ccfs also depend on the galaxy populations used and their clustering strength. only recently has an alternative approach (coil et al. 2009) allowed the comparison of the results from different studies by inferring the agn acf from the measured ccf and acf of the galaxy tracer set. the basic idea of this method, which is now frequently used (e.g., krumpe et al. 2010, 2012; mountrichas & georgakakis 2012; shen et al. 2013), is that both populations trace the same underlying dm density field. photometric redshift samples: large galaxy tracer sets with spectroscopic redshifts are not available at all redshifts. some studies therefore rely on photometric redshifts. the impact of the large uncertainties and catastrophic outliers when using photometric redshifts is commonly not considered but it can be essential. the use of the full probability density function (pdf) of the photometric redshift fit, instead of a single value for the photometric redshift, has been used in some studies (e.g., mountrichas et al. 2013). here, photometric galaxies samples are used as tracer sets to derive the ccf between agn and galaxies. each object is given a weight for the probability that it is actually located at a certain redshift based on the fit to the photometric data. figure 2: in the conceptual model of the hod approach, there are two contributions to the pairs that account for the measured correlation function. pairs of objects (black stars) can either be located within the same dmh (pink filled circles), such that their measured separation contributes to the 1-halo term (red solid line in the large dmh), or can reside in different dmhs, such that their separations (green dotted line) contribute to the 2-halo term. agn halo occupation distribution modeling: instead of deriving only mean dmh masses from the linear bias parameter, hod modeling of the correlation function allows the determination of the full distribution of agn as a function of dmh mass. the derived distribution also connects observations and simulations as it provides recipes for how to populate dmhs with observable objects. in the hod approach, the measured 2pcf is modeled as the sum of contributions from pairs within individual dmhs (fig. 2; 1-halo term) and in different dmhs (2-halo term). the superposition of both components describes the shape of the observed 2pcf better than a simple power law. in the hod description, a dmh can be populated by one central agn or galaxy and by additional objects in the same dmh, so-called satellite agn and galaxies. applying the hod approach to the 2pcf allows one to determine, e.g., the minimum dmh needed to host the object class of interest, the fraction of objects in satellites, and the number of satellites as a function of dmh mass. instead of using the derived agn acf from ccf measurements, miyaji et al. (2011) utilize the hod model directly on high precision agn vs. galaxy ccf and achieve additional constraints on the agn and galaxy co-evolution and agn physics. 75 mirko krumpe, takamitsu miyaji, alison l. coil theoretical predictions: only recently have several different theoretical models been published which try to explain the observed agn clustering with different physical approaches (e.g, fanidakis et al. 2013; hütsi et al. 2014). the key to observationally distinguish between these models are their different predictions for the clustering dependences of different agn parameters. in addition to theoretical models of the observed clustering, other very recently developed models predict the halo occupation distribution of agns at different redshifts, e.g., chatterjee et al. (2012). the major challenge presented by all of these models is the urgent need for observational constraints with higher precision than can be provided with current agn samples. in the future, progress in agn physics and agn and galaxy evolution will be achieved through a close interaction between state-ofthe-art cosmological simulations and observational constraints from high precision clustering measurements. simulations which incorporate different physical processes will lead to different predictions of the agn and galaxy large-scale clustering trends and their halo occupation distributions. observational studies will then identify the correct model and consequently the actual underlying physical processes. 5 the future of agn clustering measurements agn clustering measurements from several upcoming projects will significantly extend our knowledge of the growth of cosmic structure and will also provide a promising avenue towards new discoveries in the fields of galaxy and agn co-evolution, agn triggering, and cosmology. for example, erosita (predehl et al. 2010; launch 2015/2016) will perform several all-sky x-ray surveys. after four years the combined survey is expected to contain approximately three million agns. hetdex (hill et al. 2008; start 2015) will use an array of integral-field spectrographs to provide a total sample of ∼20,000 agns without any pre-selection over an area of ∼ 450 deg2. the sdss-iv/eboss and bigboss builds upon the sdss-iii/boss project and will use a fiber-fed spectrograph. over an area of 14,000 deg2, it will observe roughly one million qsos at 1.8 < z < 3.5. in addition to these projects, there will be other major enterprises such as lsst (lsst collaboration 2009) and pan-starrs (kaiser et al. 2002) which will detect several million agns but currently lack dedicated spectroscopic follow-up programs. in the following we will focus on erosita, as this mission will compile the largest agn sample ever observed. figure 3 shows that erosita agn detections will outnumber current galaxy samples with spectroscopic redshifts at z > 0.4. using a large number of agns that continuously cover the redshift space, will allow us (in contrast to galaxy samples) to measure the distribution of matter with high precision in the last ∼11 gyr of cosmic time. to fully exploit the erosita potential for agn clustering measurements, a massive spectroscopic follow-up program is needed. several spectroscopic multi-object programs and instruments are currently planned or are in an early construction phase (e.g., sdss-iv/spiders and 4most). figure 3: number of expected erosita agns (red) and currently available galaxies with spectroscopic redshifts (black solid line at z < 0.4 – sdss data release 7; black dotted line – primus (coil et al. 2011); black solid line at z ∼ 1 – deep2 (davis et al. 2003) and vvds (le fèvre et al. 2005)). instead of the full sky area, we consider only the expected number of erosita agns with spectroscopic redshifts from 4most over 14,000 deg2. erosita agn clustering measurements at z ∼ 0.8 − 1 will even allow for the detection of the bao signal. the feasibility of such a measurement can be estimated using the bao detection found with ∼46,000 sdss lrgs (〈z〉 = 0.35) over 3,816 square degrees of sky (0.72 h−3 gpc3) as a standard for comparison (eisenstein et al. 2005). the observed agn x-ray luminosity function (gilli et al. 2007) and the erosita sensitivity determine the number density of erosita agns. in the abovementioned redshift range, the erosita agn area density will be comparable to that of sdss lrgs at lower redshifts. therefore, the comoving volume number density of erosita agns will be five times lower than that of sdss lrgs. since erosita will conduct an all-sky survey, the increased sky area will counterbalance the lower volume density. given the signal-to-noise ratio (s/n) of the bao detection of eisenstein et al. (2005) and an assumed spectroscopic area of 14,000 deg2, we expect a ∼3σ bao detection using erosita agns only in the redshift range of z ∼ 0.8 − 1. this is consistent with 76 clustering measurements of broad-line agns: review and future kolodzig et al. (2013), who use a different approach based on the angular power spectrum for estimating the significance of a bao detection with erosita agns. with the much larger agn datasets that will exist in the future, the statistical uncertainties in clustering measurements will be significantly decreased. systematic uncertainties will then be the dominant source of uncertainty. the impact and level of different systematic uncertainties can only be carefully explored and quantified through simulations. thus far, there has not been a need for such studies because the agn samples to date are i) drawn from surveys that (with exceptions) cover a rather moderate sky area and are therefore likely to suffer from the problem of cosmic variance2 and/or ii) comprised of up to several thousand objects and are consequently poisson noise dominated. both limitations will be removed in future agn clustering measurements with the upcoming extensive agn samples covering extremely large sky areas. however, to derive reliable constraints on agn physics and cosmology, as well as to avoid any possible misinterpretations of future unprecedented high precision agn clustering measurements, we have to fully understand and be able to correctly model the impact of the systematic uncertainties. only then can we maximize the scientific return of future agn clustering measurements and have a major impact in the field of cosmology and galaxy and agn evolution. acknowledgement mk received funding from the european community’s seventh framework programme (/fp7/20072013/) under grant agreement no 229517. tm is supported by unam/papiit in104113 and conacyt 179662. alc acknowledges support from nsf career award ast-1055081. references [1] abazajian et al. 2009, apjs, 182, 543. doi:10.1088/0067-0049/182/2/543 [2] allevato et al. 2011, apj, 736, 99. doi:10.1088/0004-637x/736/2/99 [3] bolton et al. 1976, apj, 210, l1. 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[54] weinberg 2002, aspc, 283, 3. 78 http://dx.doi.org/10.1088/0067-0049/192/2/16 http://dx.doi.org/10.1086/318331 http://dx.doi.org/10.1088/0004-637x/726/2/83 http://dx.doi.org/10.1111/j.1365-2966.2011.20059.x http://dx.doi.org/10.1093/mnras/sts666 http://dx.doi.org/10.1046/j.1365-8711.2000.03779.x http://dx.doi.org/10.1046/j.1365-8711.2001.04006.x introduction understanding observed clustering properties clustering measurements why are we interested in agn clustering? agn clustering measurements: past and present recent developments the future of agn clustering measurements acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0013 acta polytechnica ctu proceedings 4:13–18, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app modelling of nuclear fuel cladding tubes corrosion miroslav cecha, ∗, martin seveceka, b a department of nuclear reactors, faculty of nuclear sciences and physical engineering, czech technical university in prague, v holesovickach 2, praha 8, czech republic b alvel a.s., opletalova 37, praha 1, czech republic ∗ corresponding author: miroslav.cech@post.cz abstract. this paper describes materials made of zirconium-based alloys used for nuclear fuel cladding fabrication. it is focused on corrosion problems their theoretical description and modeling in nuclear engineering. keywords: corrosion, zirconium alloys, nuclear fuel, cladding, modelling. 1. introduction tubes covering nuclear fuel in current light water reactors (lwr) are made of zirconium-based alloys since the very origin of nuclear power utilization. zirconiumbased alloys were first used in nuclear reactors of u.s. nuclear submarines. later, zr was adopted by fuel vendors as a suitable material for fuel cladding for commercial reactors around the world. zirconium has been chosen for its low cross section for neutron absorption, good corrosion resistance, and other outstanding thermomechanical characteristics. various degradation processes jeopardizing nuclear cladding integrity take place during reactor normal operation such as grid-to-rod fretting, debris-induced failures, crud-induced localized corrosion, waterside corrosion, and hydriding. this article is focused on water corrosion, its quantification, and theoretical description. corrosion reaction caused mainly by an interaction of nuclear fuel cladding and coolant takes place on an external surface of cladding tubes, less frequently reacts internal cladding surface with oxygen released from pelets. in case of lwrs, metal reacts with water and zirconium oxide arises: zr + 2 h2o −−→ zro2 + 2 h2. (1) hydrogen released from the reaction described above partly dissolve in coolant water and partly penetrates tubes causing hydriding of zirconium. the fraction of hydrogen released from the reaction that is locally absorbed by the cladding tube is called pickup fraction and in pwr conditions is found to be constant for particular zr-based alloys [1]. oxygen diffuses through the zirconium oxide layer and in an interface of metal and oxide causes additional oxidation. the density of the zro2 is smaller than zirconium alloy density. the difference in density and different thermal expansion of materilas is the primary cause of tension, internal stresses, and strains in cladding tubes. moreover, thermal conductivity λ of zro2 is much smaller than that of zirconium based alloys causing the zirconium dioxide layer to decrease the heat transfer from the fuel pellet over the cladding to the coolant and consequently increase maximal temperature in the fuel pellet. the exact physical parameters depend on temperature and models of thermal conductivity of zirconium dioxide. they are in details described in [2] and [3]. for example, the thermal conductivity of e110 alloy is about 18 w/mk at the temperature of 300 ◦c. the thermal conductivity of the zro2 for the same temperature is only about 2 w/mk. the bulk density of the zircaloy-4 alloy and zirconium dioxide at the same temperature is 6.5 and 5.6 kg/m3 respectively. when the zirconium dioxide layer thickness exceeds about 100 µm (zircaloy-4), it might crack and it is washed away by the streaming coolant which can lead to cold spots, additional oxidation, hydridation and later cladding breach [4]. oxidation takes place also on the internal surface of fuel cladding, where metal reacts with oxygen released from the fuel pellets where fission takes place. when a high burn-up is reached, a bonding layer consisting of zro2, uo2, and fission products appears. this layer might be the cause of a firm connection between the fuel cladding and pellets. 2. zirconium based alloys zirconium-based alloys has been used as the nuclear fuel cladding since 1950s [6]. there were originally two main groups of zr-based alloys developed: (1.) zirconium-tin and iron-based alloys (originally developed in the u.s.) (2.) zirconium-niobium based alloys (originally developed in former ussr) during the evolution of the nuclear fuel, fuel vendors and research organizations developed dozens of concepts of fuel cladding alloys. however, the two main groups remained as can be seen in figure 1. different cladding concepts can be based on the figure 13 http://dx.doi.org/10.14311/ap.2016.4.0013 http://ojs.cvut.cz/ojs/index.php/app miroslav cech, martin sevecek acta polytechnica ctu proceedings figure 1. nuclear fuel cladding alloys developed for usage in pwr reactors [5]. there were two main groups of cladding alloys developed: zr-sn (right branch) and zr-nb (left branch) based alloys with dozens other concepts depending on the alloying elements introduced around the world. divided into four development directions depending on their composition and development history. each alloy has a different corrosion characteristics depending on the alloying elements but also on manufacturing process and reactor type. 2.1. zirconium-tin alloys a well-known family of alloys called zircaloy was developed under a westinghouse-led program. an alloy first developed is called zircaloy-1. this alloy was replaced by the innovative zircaloy-2, which is still in use in the bwr reactors. after abandoning of the zircaloy-3 alloy development and utilization due to metallurgical processing issues, concerns with the high hydrogen pickup fraction exhibited by the zircaloy-2 alloy led to the development of the zircaloy4 alloy. nickel was substituted by iron in this alloy and it was used from the 1950s to 1990s. 2.2. zirconium-niobium alloys an alloy called e110 has been used by the russian nuclear industry for nuclear fuel cladding fabrication for vver reactors fuels. similar alloys made of zirconium-doped by niobium were used from the origin of the nuclear power utilization in the ussr. recently, westinghouse replaced zircaloy-4 alloy in most of their nuclear fuel for pwrs by the zirlo alloy. in bwr reactors, zircaloy-2 is still in use. zirlo is doped with niobium and is similar to the russian alloy called e635. french alloy called m5 is a zirconium-based alloy containing 1 % of niobium with oxygen. the m5 alloy is produced by the french company areva since 1990s. generally, alloys doped with niobium such as e110, m5, and zirlo have a higher corrosion resistance than alloys from the zircaloy family. summary of composition of currently widely employed alloys is presented in table 1 [7]. 3. corrosion models the growth and development of the cladding corrosion layer for each zirconium-based alloy is usually expressed by corrosion models. 3.1. garzarolli models models developed by garzarolli et al. [8] are adopted for describing the thickness of the corrosion layer of cladding tubes made of zircaloy-4 , m5, and zirlo alloys under pwr conditions. these models generally based on arrhenius law divide a growth of corrosion layer into two phases: (1.) the first phase continues until a transition thickness of oxide layer str is reached. the rate of corrosion layer growth is expressed by the cubic rate law – equation (2) (original units and the temperature of oxide-metal interface is used) [2, 3]. (2.) after reaching an alloy-specific transition thickness of oxide layer str, second phase quantified by a linear differential equation (3) taking into account also fast neutron flux φ is adopted. there are other transition points observed during the corrosion process. however, only the first transition is significant for the general progress of the corrosion kinetics. for modelling purposes the use of only one transition is satisfactory. model is defined for temperature range of 250–400 ◦c. ds3 dt = a s2 exp ( − q1 rt ) , (2) ds dt = (c0 + u(mφ)p ) exp ( − q2 rt ) , (3) where s . . . oxide layer thickness [µm], t . . . temperature [◦c], φ . . . neutron flux [neutrons/m2s], a = 6.3 × 109 m3/day, r = 1.98 cal/mol k, c0 = 8.04 × 107 µm/day, m = 7.46 × 10−15 cm2s/n, p = 0.24, u = 2.59 × 108 µm/day. values of constants q1, q2 and transition thickness of oxide layer str are alloy-dependent and are summarized in table 2 for four widely used alloys which were subject to many studies. in literature, there were modifications of correlation models (2) and (3) defined. modified models differ by values of constants and consider other physical phenomena neglected in the presented models. these models are implemented for example in femaxi and frapcon fuel performance codes. 14 vol. 4/2016 modelling of cladding tubes corrosion alloying element zircaloy-4 zirlo m5 e-110 e-635 sn 1.2–1.7 % 0.96–0.98 % 1.25–1.30 % fe 0.18–0.24 % 0.094–0.105 % <500 ppm 0.30–0.35 % cr 0.07–0.13 % 79–83 ppm <200 ppm nb 1.02–1.04 % 0.8–1.2 % 0.9–1.1 % 1.0 % ni 0.01 % n <65 ppm 22–30 ppm <60 ppm 30–40 ppm c 150–400 ppm 60–80 ppm <200 ppm < 200 ppm o 900–1400 ppm 900–1200 ppm 0.11–0.17 % <1000 ppm 0.07 % table 1. composition of zirconium-based alloys widely used around the world in pwrs as materials for cladding tubes fabrication [7]. q1[cal/mol] q2[cal/mol] str zircaloy-4 32289 27354 2 µm m5 27446 29816 7 µm zirlo 32289 27080 2 µm opt.zirlo 32289 27354 2 µm table 2. values of corrosion model’s constants used in corrosion models (2) and (3) as defined by [3]. figure 2. zirconium dioxide layer development from the point of reaching first transition thickness str corresponding to the transition time ttr. axis x represents time in days after the moment, when transition thickness was reached. the temperature of 320 ◦c and neutron flux of 1 × 1015 neutrons/m2s were chosen in this model situation. 3.2. three-phase model another model for describing of zircaloy-4 corrosion layer thickness in pwr conditions divides its evolution into three phases instead of two. the purpose is a faster oxide thickness growth after the second transition thickness is reached [7]. the model was developed by experimental data fitting and is more precise for higher values of oxide thickness than model developed by garzarolli et al. [8]. oxide layer growth during the first phase can be calculated by the following expression [6] ds3 dt = 2.187 × 10−13 exp ( − 135188 rt ) . (4) the first transition thickness is the same as in the previous model – 2 µm for zircaloy-4. afterward, different formula is used instead of equation (3) ds dt = ( 9.31 × 10−4 + 2.75 × 10−3 ( φ 5.24 × 1018 )0.24) · exp ( − 114526 rt ) . (5) after reaching the second transition thickness – 35 µm, following equation is used ds dt = ( 9.31 × 10−4 + 2.75 × 10−3 ( φ 5.24 · 1018 )0.24 ) · 1.8 exp ( − 114526 rt ) . (6) 3.3. e110 corrosion model a model describing the corrosion layer growth of e110 alloy in vver conditions was developed by fitting experimental data at the russian a.a. bochvar hightechnology scientific research institute for inorganic materials. for the e110 alloy following relation was derived [9] ds dt = 40 exp ( − 5147 t ) . (7) model was derived byl data base on experiments, which took place in temperature range of 320–360 ◦c. this model considers only one phase of corrosion layer development, transition thickness is disregarded. comparison between corrosion layer growth of zircaloy-4 and e110 is plotted in figure 3. figure 3. comparison of the zirconium oxide layer thickness of zircaloy-4 and e110 alloys in conditions of 320 ◦c and neutron flux of 1e15 neutrons/m2s. the corrosion growth of e110 is considerably lower in comparison with the zircaloy-4 alloy. accelerating 15 miroslav cech, martin sevecek acta polytechnica ctu proceedings growth can be seen for the zircaloy-4 three-phase model which is not the case for the e110 alloy. for that reason, the neglecting of transitions in case of e110 is justified. if the behavior of nuclear fuel during loca, ria or other design basis accidents is calculated by fuel performance codes like fraptran or transuranus, the thickness of oxide layer is an initial condition which strongly influences a progress of the accident and its consequences. 4. corrosion in loca conditions large break loss of coolant accident (lbloca) is the maximal design basis accident of pwrs of second and third generation. during this accident, a temperature of the whole fuel rod including pellets quickly rises due to limited cooling conditions. the high temperature of fuel pellets leads to high-temperature gradients, stresses, and cracking of pellets. rapid release of fission gases from the fuel can be observed and the internal pressure of filling gas rises. the high temperature of cladding together with the high internal pressure can be a cause of a deformation, ballooning, or bursting of the cladding. this geometry changes can limit the coolant flow and further reduce the heat transfer from the fuel rods to the coolant. construction of reactor, design of the nuclear fuel and its properties must ensure that acceptance limits for loca accidents will not be violated: (1.) a peak cladding temperature can nowhere exceed 1204 ◦c (2.) sufficient fuel rod strength upon quench taking into account an additional mechanical load (maintain post-quench ductility) (3.) fraction of zirconium reacted into oxide cannot exceed 1 % (due to hydrogen production) (4.) melting temperature of fuel can not be reached in any place of the reactor core 4.1. corrosion models in loca conditions to develop corrosion models in loca conditions, it is necessary to measure experimental data in similar conditions. experiments are done at high-temperature steam (800–1200 ◦c). experiments with as-received cladding tubes and as well as with cladding tubes with corrosion layer has been performed. preoxidation of experimental samples ensures that simulation will be performed in conditions which are similar to the real loca accident conditions with operating fuel. a model of high-temperature corrosion of sponge based e110 alloy was developed at the ujp in the czech republic and is based on its experimental data [10]. experiments cover a wide range of conditions (temperature 600–1300 ◦c and 0–480 minute long exposition). these wide conditions enable to use the model in various conditions and for various scenario for e110 alloy. no. t [◦c] process 1 600–750 phases α + β transformation 2 750–950 formation of β phase 3 950–1100 delated transformation, tetragonal oxide 4 1100–1300 tetragonal oxide table 3. physical processes taking place in zirconium e110 alloy in different temperature intervals during loca accident conditions. this model describes a mass growth of oxide as defined in [10] ∆g = a exp ( e t ) tn = ktn, (8) where ∆g . . . mass growth [mg/dm2], a,e,k . . . fitting parameters, t . . . time [s], n . . . kinetic constant. figure 4. development of the parameter n with temperature as defined by [10]. figure 5. development of the parameter k with temperature as defined by [10]. for n and k parameters were derived following formulas n = 0.4 for t < 768.4 ◦c (9) n = 2.609 − 4.898 × 10−3(t − 273.15) + 2.633 × 10−6(t − 273.15)2 for t < 960.3 ◦c (10) n = 1.202 × 10−3(t − 273.15) − 0.8208 for t < 1098.9 ◦c (11) n = 0.5 for t > 1098.9 ◦c (12) 16 vol. 4/2016 modelling of cladding tubes corrosion figure 6. power history of rods bsm-25 and bk365. figure 7. oxide layer thickness of bsm-25 and bk365 rods. k = 85265.6 exp ( − 9875.59 t ) for t < 934.1 ◦c (13) k = 1072.21 exp ( − 4592.6 t ) for t < 1054.5 ◦c (14) k = 33.33 for t < 1098.0 ◦c (15) k = 96482.3 exp ( − 10913.1 t ) for t > 1098.0 ◦c (16) temperature intervals of equations (9)–(16) approximately correspond to physical processes, which take place in the cladding material during loca accident. these processes are described in table 3. this model well describes a corrosion kinetic for all ranges of temperature covered by experiment. the value of parameter n is 1/2 for high-temperature corrosion and 1/3 for middle-temperature corrosion [7]. these values are the same as in other used models. another model for n also very well describes the n temperature reliance. constant k equals approximately equals to 0 and increase with temperature to about 90 at 1300 ◦c. dividing model into four ranges where different formulas are used brings a good agreement of the model with experimental data. a comparison between this model and data can be found in [10]. 5. model of corrosion in femaxi a calculation of oxide layer development in the femaxi-6 code has been performed for two fuel rods: bsm-25, and bk365. these rods were irradiated in the br-3 reactor and reached burn-up of 66 and 52 mwd/kghm. rods were irradiated within the high burnup effect program in the br-3 reactor, cladding was made of zircaloy-4 alloy, coolant inlet temperature was 255 ◦c. a model originally developed by garzarolli et al. [8] (equations (2) and (3)) was used for the zircaloy-4 alloy corrosion modeling. the two power histories used as a model input are plotted in figure 6. corresponding oxide layer thickness growth is shown in figure 7. for both tested fuel rods, an average thickness and the maximal oxide layer thickness are shown. maximal thickness of the corrosion layer is about two times higher than the average value. maximal thickness was reached in the middle of the fuel rod which does 17 miroslav cech, martin sevecek acta polytechnica ctu proceedings not correspond to the case in commercial power reactors. the oxide layer growth is strongly dependent on the temperature, the higher the temperature is the faster the growth. therefore, the maximal thickness of the zirconium oxide layer in commercial power reactors is at the top of the fuel rod where the coolant temperature is highest. figure 7 shows, that higher linear heat rate causes faster oxide layer creation. the graph also shows that the first phase of oxide layer growth is independent of fast neutron flux. small differences are caused by higher temperature. in the later phase of fuel rod’s irradiation, there is a clear dependance of corrosion layer thickness on fast neutron flux and linear heat rate. when bsm-25 rod was operating in high heat rate condition, layer growth was much faster than in the case of smaller heat rate of the second fuel rod. higher heat rate (and corresponding temperature) causes the bigger corrosion layer, even when the burnup of the bsm-25 rod is lower. 6. conclusion this article describes models quantifying corrosion of nuclear fuel cladding tubes made of zirconium-based alloys widely used in nuclear industry in nominal conditions and loca accident conditions. models used for calculating of an oxide layer thickness in normal operation conditions for the widely used alloys zircaloy-4, zirlo, m5 and e110 are presented and compared. a model for corrosion and high-temperature oxidation in loca conditions is described and reliance of particular parameters used in models are shown in graphs. a corrosion model for nominal conditions was applied in the femaxi-6 code to calculate corrosion of fuel rods bk363 and bsm-25 tested in the br-3 reactor. the relations of burnup, linear heat rate, and corrosion layer thickness growth is illustrated in the example. results show faster oxide growth in case of bsm-25 rod after reaching the transition thickness for zircaloy-4 alloy, because this rod was operated at higher linear heat rate. acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs ohk4-008/16. references [1] k. geelhood, w. luscher. frapcon-3.5: integral assessment. progress in nuclear energy 2014. [2] m. suzuki, h. saitou, n. g. k. k. kikō. light water reactor fuel analysis code: femaxi-6 (ver. 1): detailed structure and user’s manual. japan atomic energy research institute, 1997. [3] w. g. luscher, k. j. geelhood, et al. material property correlations: comparisons between frapcon-3.4, fraptran 1.4, and matpro. us nuclear regulatory commission, office of nuclear regulatory research, 2011. [4] p. billot, b. cox, k. ishigure, et al. corrosion of zirconium alloys in nuclear power plants. in tecdoc684. international atomic energy agency (iaea), 1993. [5] l. hallstadius, s. johnson, e. lahoda. cladding for high performance fuel. progress in nuclear energy 57:71–76, 2012. [6] a. t. motta, a. couet, r. j. comstock. corrosion of zirconium alloys used for nuclear fuel cladding. annual review of materials research 45:311–343, 2015. [7] d. kobylka. termofyzikální vlastnosti pokrytí nových typů paliva, jejich implementace v kódu femaxi-6 a testování. tech. rep. 14117/2012/02kob, újv řež, a.s., 2012. [8] f. garzarolli, d. jorde, r. manzel, et al. review of pwr fuel rod waterside corrosion behavior. tech. rep., kraftwerk union ag, erlangen (germany, fr); combustion engineering, inc., windsor, ct (usa), 1980. [9] v. konkov, m. sablin, t. khokhunova, et al. assessment of e110 and e635 alloy corrosion behavior in vver-1200 reactors. jsc vniinm 2009. [10] j. krejčí. oxidace palivového pokrytí během havárie loca. bezpečnost jaderné energie 45:311–343, 2015. 18 acta polytechnica ctu proceedings 4:13–18, 2016 1 introduction 2 zirconium based alloys 2.1 zirconium-tin alloys 2.2 zirconium-niobium alloys 3 corrosion models 3.1 garzarolli models 3.2 three-phase model 3.3 e110 corrosion model 4 corrosion in loca conditions 4.1 corrosion models in loca conditions 5 model of corrosion in femaxi 6 conclusion acknowledgements references 273 acta polytechnica ctu proceedings 2(1): 273–276, 2015 273 doi: 10.14311/app.2015.02.0273 the ubv color evolution of classical and symbiotic novae i. hachisu1, m. kato2 1department of earth science and astronomy, college of arts and sciences, the university of tokyo, komaba, meguroku, tokyo 153-8902, japan 2department of astronomy, keio university, hiyoshi, kouhoku-ku, yokohama 223-8521, japan corresponding author: hachisu@ea.c.u-tokyo.ac.jp abstract we identified a general course of classical nova outbursts in the b − v vs. u − b diagram. it has been reported that novae show spectra similar to a–f supergiants near optical light maximum. however, they do not follow the supergiant sequence in the color-color diagram, neither the blackbody nor the main-sequence sequence. instead, we found that novae evolve along a new sequence in the pre-maximum and near-maximum phases, which we call the nova-giant sequence. this sequence is parallel to but ∆(u − b) ≈ −0.2 mag bluer than the supergiant sequence. after optical maximum, its color quickly evolves back blueward along the same nova-giant sequence and reaches the point of free-free emission (b − v = −0.03, u − b = −0.97) and stays there for a while, which is coincident with the intersection of the blackbody sequence and the nova-giant sequence. then the color evolves leftward (blueward in b − v but almost constant in u − b) due mainly to development of strong emission lines. this is the general course of nova outbursts in the color-color diagram, which is deduced from eight well-observed novae including various speed classes. for a nova with unknown extinction, we can determine a reliable value of the color excess by matching the observed track of the target nova with this general course. this is a new and convenient method for obtaining color excesses of classical novae. using this method, we redetermined the color excesses of nineteen well-observed novae. keywords: cataclysmic variables novae individual: fh ser, pu vul, pw vul, v1500 cyg, v723 cas. 1 introduction a classical nova is a thermonuclear runaway event on a mass-accreting white dwarf (wd) in a binary. despite of their overall similarity, optical light curves of novae have a wide variety of timescales and shapes. recently, hachisu & kato (2006) found that, in terms of free-free emission, the optical and infrared (ir) light curves of several novae follow a similar decline law. moreover, the time-normalized light curves were found to be independent of the wd mass, chemical composition of ejecta, and wavelength. they called it the universal decline law. evolution of colors is another challenging subject that attracts many researchers, who attempted to find a general behavior among various types of novae. for example, duerbeck & seitter (1979) discussed that color evolutions of six novae are remarkably similar in the intrinsic (b−v )0 versus (u−b)0 color-color diagram independently of their different nova speed classes. van den bergh & younger (1987) derived general trends of color evolutions in nova light curves, i.e., (b − v )0 = 0.23 ± 0.06 at optical maximum and (b − v )0 = −0.02 ± 0.04 at t2, where tm (m = 2 or 3) is the days during which a nova decays by mth magnitude from its optical maximum and (b −v )0 is the intrinsic b − v color of the nova. these two relations, however, often show large deviations from the values obtained by other methods. miroshnichenko (1988) found a stabilization stage of novae in the b−v and u − b color evolutions soon after optical maximum and that this stage showed a general trend of (b − v )0 = −0.11 ± 0.02. he derived extinctions, e(b−v ), of 23 novae assuming that all novae have the same intrinsic (b−v )0 color at the stabilization stage, i.e., e(b−v ) = (b−v )ss−(b−v )0 = (b−v )ss +0.11, where (b−v )ss is the observed b−v color at the stabilization stage. this method looks powerful but sometimes results in a large difference from the true value. according to hachisu & kato’s (2006) universal decline law, optical fluxes in the ubv bands could be dominated by free-free emission. if it is the case, its color is simply estimated to be (b − v )0 = 0.13 and (u −b)0 = −0.82 for the optically thin free-free emission (fν ∝ ν0), or to be (b − v )0 = −0.03 and (u −b)0 = −0.97 for the optically thick free-free emission (fν ∝ ν2/3, see, e.g., wright & barlow 1975), where fν is the flux at the frequency ν. the latter value of (b−v )0 = −0.03 is close to (b−v )0 = −0.02±0.04 273 http://dx.doi.org/10.14311/app.2015.02.0273 i. hachisu, m. kato at t2 derived by van den bergh & younger (1987). however, many novae do not keep these pivot points but further evolves blueward. these different trends of nova color evolutions may represent different sides of the true color evolution which we do not yet fully understand observationally or theoretically. the aim of this presentation is to find a general path of nova color evolutions as duerbeck & seitter (1979) tried to find about thirty years ago. first we examine the color evolutions of slow and fast novae, fh ser, pw vul, v1500 cyg, v1668 cyg, v1974 cyg, pu vul, hr del, and v723 cas. we found that the tracks of color-color evolutions of these slow/fast novae are common. then we redetermined the reddening of nineteen novae by fitting the color evolution of a target nova with the common path in the color-color diagram. thus, we propose a new method for estimating the color excess. figure 1: light and color curves of fh ser 2 nova-giant sequence in the color-color diagram the first example of our analysis is the moderately fast nova fh ser. the v , (b − v )0, and (u − b)0 light curves of fh ser are plotted in figure 1. fh ser showed a gradual optical decay with t2 = 41 and t3 = 62 days followed by a sudden drop of the brightness due to dust shell formation about 2.8 mag below the optical maximum. the (b−v )0 vs. (u −b)0 color-color evolution is plotted in figure 2 (see hachisu & kato 2014 for more detail). the reddening toward fh ser was obtained by kodaira (1970) to be e(b − v ) = 0.6 from the mmrd relation and interstellar reddening relation. della valle et al. (1997) obtained e(b − v ) = 0.82 from the color at optical maximum, i.e., e(b − v ) = (b − v )max − (b − v )0,max = 1.05 − 0.23 = 0.82, e(b −v ) = 0.61 from the line ratio of hα/hβ = 5.7, and e(b −v ) = 0.5 from the equivalent width of na i λ5890. then, della valle et al. (1997) adopted an averaged value of e(b −v ) = 0.64 ± 0.16. since kodaira’s e(b−v ) = 0.60 and della valle et al.’s e(b−v ) = 0.61 from the line ratio of hα/hβ are coincident with each other, we use e(b−v ) = 0.60 in the present paper and confirm below that this value is reasonable. figure 2: color-color evolution of fh ser in figure 2 we plot only the data before the dust blackout started about 70 days after the outburst. we also depict three known sequences, the blackbody, supergiant, and main-sequence sequence, the data of which are taken from allen (1976). we also added a point, i.e., optically thick free-free emission spectra (open diamond denoted by f) of fν ∝ ν2/3. we have frequently seen in literature such a statement that nova spectra near maximum are similar to those of a–f type supergiants. however, the track of fh ser in the colorcolor diagram does not follow the supergiant sequence as clearly shown in figure 2. the shape of fh ser track is very similar to that of the supergiant sequence but 274 the ubv color evolution of classical and symbiotic novae it locates about ∆(u −b) ≈ −0.2 mag bluer than the supergiant sequence. therefore, we are forced to define a new sequence based on the data of fh ser, which is designated by points a, b, c, d, and f from redder to bluer. these points correspond to epochs in figure 1. in what follows, we will see that many novae evolve along this sequence when their photospheric spectra are similar to those of a–f type supergiants. therefore, we call this track “the nova-giant sequence” after the supergiant sequence. 3 common paths of novae in the color-color diagram we have also studied color-color evolutions of the fast and slow novae, pw vul, v1500 cyg, v1668 cyg, v1974 cyg, pu vul, hr del, and v723 cas, in the (b−v )0 (u−b)0 diagram and showed that these well observed novae follow very similar evolutionary courses in the intrinsic color-color diagram as shown in figure 3. we finally specified four templates, (1) the moderately fast nova fh ser, a proto-type of the nova-giant sequence, (2) the slow nova pw vul and the very fast nova v1500 cyg, (3) the fast novae v1668 cyg and v1974 cyg, and (4) the symbiotic nova pu vul, the very slow novae v723 cas and hr del. the tracks are characterized with several specified points (a, b, c, d, f, 0, 1, ..., 5, 4’, 5’, 4”, and 5”). figure 4(a) shows templates of these eight well observed novae. figure 3: several typical tracks of novae in the colorcolor diagram since these well observed eight novae follow the templates in figure 4(a), we expect other novae also follow the same paths. in other words, if all novae follow this template, we are able to determine the color excess of a target nova by directly comparing its track in the colorcolor diagram with our general track in figure 4(a). figure 4: color-color evolution of (a) template 8 novae, (b) rs oph, (c) v446 her, and (d) v533 her. figure 5: (a) t pyx, (b) lv vul, (c) iv cep, and (d) nq vul. 275 i. hachisu, m. kato 4 estimates of extinctions toward various novae we are able to determine color excesses of various novae using our general color-evolution tracks found in this work. we collected novae as many as possible from literature that have sufficient data points (usually more than ten). we assume that all novae follow the colorcolor evolution tracks in figure 4(a). in order to obtain e(b − v ), we change e(b − v ) by steps of 0.05 to fit the observed track of a target nova with our general tracks in figure 4(a). figure 6: (a) v1370 aql, (b) gq mus, (c) qu vul, and (d) os and. figure 7: (a) qv vul, (b) v443 sct, (c) v1419 aql, and (d) v705 cas. figure 8: (a) v382 vel, (b) v2274 cyg, (c) v475 sct, and (d) v5114 sgr. figures 4, 5, 6, 7, and 8 show results of our analysis for 19 novae. corresponding e(b−v ) excess is denoted in each panel (see hachisu & kato 2014, in detail). references [1] allen, c. w. 1976, astrophysical quantities (3rd ed.; london: athlone) [2] della valle, m., gilmozzi, r., bianchini, a., & esenoglu, h. 1997, a&ap, 325, 1151 [3] duerbeck, h. w., & seitter, w. c. 1979, a&ap, 75, 297 [4] hachisu, i., & kato, m. 2006, apjs, 167, 59 doi:10.1086/508063 [5] hachisu, i., & kato, m. 2014, apj, 785, 97 doi:10.1088/0004-637x/785/2/97 [6] kodaira, k. 1970, pasj, 22, 447 [7] miroshnichenko, a. s. 1988, soviet astronomy, 32, 298 [8] van den bergh, s., & younger, p. f. 1987, a&ap suppl., 70, 125 [9] wright, a. e., & barlow, m. j. 1975, mnras, 70, 41 doi:10.1093/mnras/170.1.41 276 http://dx.doi.org/10.1086/508063 http://dx.doi.org/10.1088/0004-637x/785/2/97 http://dx.doi.org/10.1093/mnras/170.1.41 introduction nova-giant sequence in the color-color diagram common paths of novae in the color-color diagram estimates of extinctions toward various novae 293 acta polytechnica ctu proceedings 1(1): 293–297, 2014 293 doi: 10.14311/app.2014.01.0293 design and tests of the hard x-ray polarimeter x-calibur m. beilicke1, r. cowsik1, p. dowkontt1, q. guo1, f. kislat1, s. barthelmy2, t. okajima2, j.w. mitchell2, j. schnittman2, b. zeiger2, g. de geronimo3, m.g. baring4, a. bodaghee5, t. miyazawa6, k.d. finkelstein7, h. krawczynski1 1department of physics and mcdonnell center for the space sciences, washington university, st. louis, mo, usa 2goddard space flight center, md, usa 3brookhaven national lab, ny 4rice university, tx, usa 5uc berkeley, ca 6nagoya university, japan 7cornell university, ny corresponding author: beilicke@physics.wustl.edu abstract x-ray polarimetry promises to give qualitatively new information about high-energy astrophysical sources, such as binary black hole systems, micro-quasars, active galactic nuclei, and gamma-ray bursts. we designed, built and tested a hard x-ray polarimeter, x-calibur, to be used in the focal plane of the infocus grazing incidence hard x-ray telescope. x-calibur combines a low-z compton scatterer with a czt detector assembly to measure the polarization of 20 − 60 kev x-rays making use of the fact that polarized photons compton scatter preferentially perpendicular to the electric field orientation; in principal, a similar space-borne experiment could be operated in the 5−100 kev regime. x-calibur achieves a high detection efficiency of order unity. keywords: x-rays polarization black hole infocus x-calibur. 1 introduction motivation. spectral and morphological studies in the x-ray energy band have become established tools to study the non-thermal emission processes of various astrophysical sources. however, many of the regions of interest (black hole vicinities, jet formation zones, etc.) are too small to be spatially resolved with current and future instruments. spectro-polarimetric x-ray observations are capable of providing additional information – namely the fraction and orientation of linear polarization – and would help to constrain different emission models [1, 2] of sources with compact emission regions and high x-ray fluxes such as mass-accreting black holes (bhs) and neutron stars. so far, only a few missions have successfully measured polarization in the soft (oso-8 [3]) and hard (integral [4]) x-ray energy regime. the crab nebula is the only source for which the polarization of the x-ray emission has been established with a high level of confidence [3, 4]. the source exhibits a polarization fraction of 20% at energies of 2.6−5.2 kev (direction angle of 30◦ with respect to the x-ray jet) [3] and 46% ± 10% above 100 kev (direction aligned with the x-ray jet observed in the nebula). integral observations of the x-ray binary cygnus x-1 indicate a high fraction of polarization in the hard x-ray/gamma-ray bands [5]. model predictions of polarization fraction for various source types lie slightly below the sensitivity of the past oso-8 mission, making future, more sensitive polarimeter missions particularly interesting. future missions. as polarimetry was not the main objective of the integral mission, the results are plagued by systematic uncertainties, and there are currently no other missions in orbit that are capable of making sensitive x-ray polarimetric observations. the situation could be changed by a mission like the gravity and extreme magnetism smex (gems) [6] – using two wolter-type x-ray mirrors to focus 2 − 10 kev photons onto photo-effect polarimeters. for higher energies e > 5 kev x-ray polarimeter designs can make use of the compton effect. the soft gamma-ray imager on astro-h [7] (launch scheduled for 2015) will have capabilities of measuring polarization, but the results may suffer from similar systematic uncertainties as the ones 293 http://dx.doi.org/10.14311/app.2014.01.0293 m. beilicke et al. figure 1: left: incoming x-rays are compton-scattered (scintillator rod, optical axis) and subsequently photoabsorbed in one of the surrounding czt detectors. middle left: ’exploded’ view of the polarimeter: four sides of detector columns surround the central scintillator rod. middle right: polarimeter with csi shield, electronic readout and azimuthal rotation bearing. right: the infocus balloon gondola. measured by integral. the czt imager on astrosat (http://astrosat.iucaa.in/) is expected to be used for polarization measurements in the 150 − 250 kev band. the hard x-ray polarimeter x-calibur discussed in this paper has the potential to cover the energy range above 10 kev. furthermore, x-calibur combines a high detection efficiency with a low level of background and has well-controlled systematic errors. these features make it a particularly useful instrument for astronomical x-ray polarimetry. scientific potential. polarization measurements are of general interest as tests of non-thermal emission processes in the universe. synchrotron emission, for example, will result in linearly polarized photons with their electric fields oriented perpendicular to the magnetic field lines (projected) and the observed polarization map in the x-rays can therefore be used to trace the magnetic field structure of the source. an observed polarization fraction close to theoretical limits can be interpreted as an indication of a highly ordered magnetic field since non-uniformities in the magnetic field will reduce the fraction of polarization. the polarized synchrotron photons can be inverse-compton (ic) scattered by relativistic electrons – weakening the fraction of polarization (but not erasing it) and imprinting a scattering angle dependence [8] to the observed fraction of polarization. such ic signals usually (but not always) appear in hard gamma-rays, where polarimetry is difficult, due to multiple scattering in pair production detectors. another important mechanism for polarizing photons is thomson scattering which creates a polarization perpendicular to the scattering plane. curvature radiation is polarized, as well. for more details on the scientific prospects see for example [1] and references therein. addressing these science goals requires spectro-polarimetric observations over the broadest possible energy range. 2 x-calibur design the conceptual design of the x-calibur polarimeter is illustrated in figure 1. a low-z scintillator (ρ ≈ 1 g/cm3) is used as compton-scatterer: the cross section for photoelectric absorption and compton scattering are equal around 15 kev (0.26 cm2/g). at 20 kev the cross section of the photoelectric absorption already drops to 0.1 cm2/g and can be neglected as compared to the compton scattering for higher energies. the mean free path for the compton scattering is ≈ 4 cm so that the length of the scintillator (14 cm) covers ' 3.5 path lengths leading to a p ' 90% probability for absorption in the energy regime of 20 − 60 kev. for sufficiently energetic photons, the compton interaction produces a measurable scintillator signal which is read by a pmt. the scattered x-rays are photo-absorbed in surrounding rings of high-z cadmium zinc telluride (czt) detectors. this combination of scatterer/absorber leads to a high fraction of unambiguously detected comp294 design and tests of the hard x-ray polarimeter x-calibur ton events. linearly polarized x-rays will preferably compton-scatter perpendicular to their e field vector – resulting in a modulation of the azimuthal event distribution. the czt detectors were ordered from different companies (endicott interconnect, quikpak/redlen, creative electron). each detector (2 × 2 cm2) is contacted with a 64-pixel anode grid (2.5 mm pixel pitch) and a monolithic cathode facing the scintillator rod. two detector thicknesses (2 mm and 5 mm) will be used. each czt detector is permanently bonded (anode side) to a ceramic chip carrier which is plugged into the electronic readout board. each czt detector is read out by two digitizer boards (32 channel asic developed by g. de geronimo (bnl) and e. wulf (nrl) [9] and a 12-bit adc). the readout noise is as low as 2.5 kev fwhm. 16 digitizer boards (8 czt detectors) are read out by one harvester board transmitting the data to a pc-104 computer with a rate of 6.25 mbits/s. xcalibur comprises 2048 data channels. four detector units form a ’ring’ surrounding the scintillator slab. the scintillator ej-200 is read by a hamamatsu r7600u-200 pmt. the (optional) pmt trigger information allows to effectively select scintillator/czt events from the data, which represent likely compton-scattering candidates. the polarimeter and the front-end readout electronics will be located inside an active csi(na) anticoincidence shield with 5 cm thickness and a passive tungsten shield at the top (fig. 1, middle right) to suppress charged and neutral particle backgrounds. the x-calibur polarimeter will be flown in a pressurized vessel located in the focal plane of the infocus1 x-ray telescope [10] (fig. 1, right). the total mass of the gondola and the x-ray telescope will be 1, 400 kg. a wolter grazing incidence mirror focuses the x-rays on the polarimeter. the x-calibur scintillator rod will be aligned parallel to the optical axis of the infocus x-ray telescope. the focal length is ∼ 8 m and the field of view (fwhm) is 10 arcmin. the design of infocus allows for very stable pointing of the telescopes to <1 arcmin as the focus of the x-ray telescope moves across the sky. in order to reduce the systematic uncertainties of the polarization measurements (including biases generated by the active shield, a possible pitch of the polarimeter with respect to the x-ray telescope, etc.), the polarimeter and the active shield will be rotated around the optical axis (∼ 3 rpm) using a ring bearing (see fig. 1, middle right). a counterrotating mass will be used to cancel the net-angular momentum of the polarimeter assembly during the flight. power will be provided to the polarimeter by sliding contacts and communication will be done via a wireless network. the data will be stored to solid state drives and will be down-linked to the ground. the advantages of the x-calibur/infocus design are (i) a high detection efficiency by using more than 80% of photons impinging on the polarimeter, (ii) low background due to the usage of a focusing optics instead of large detector volumes, and (iii) minimization and control of systematic effects and achievement of a corresponding quantitative estimate thereof. detailed simulations were performed using the geant4 package [11]. an x-calibur balloon flight in the focal plane of the infocus mirror assembly was assumed. for a crab-like source the simulations predict an event rate of 1.1(3.2) hz with (without) requiring a triggered scintillator coincidence detected by the pmt. the x-calibur modulation factor is µ = 0.52 for a 100% polarized beam. the mdp in the 20 − 60 kev energy range will be 4% assuming 5.6 hr of on-source observations of a crab-like source combined with a 1.4 hr background observation of an adjacent empty field. more detailed simulations are discussed in guo et al. (2010) [12]. czt energy [kev] 20 40 60 80 /s ] 2 f lu x[ 1 /c m -310 -210 -110 1 ring1 ring2 ring3 figure 2: energy spectra after compton scattering of a 40/80 kev x-ray beam. vertical lines: incoming beam energy (solid), energy after 90◦ scattering (dashed), and energy after 180◦ scattering (dotted). 3 performance measurements using funding from washington university’s mcdonnell center for the space sciences and nasa grant nnx12ad51g, a flight-ready version of the x-calibur polarimeter was assembled and tested. for energy calibration of the czt detectors a eu152 source is placed inside the individual rings. the calibration data also allow to quantify the energy resolution and threshold on a pixel-by-pixel basis. we find an average resolution of 4.2 kev fwhm at 40 kev. in turn, we completed the full assembly and studied the x-calibur respsonse to incoming x-ray beams. in order to measure the response to a polarized xray beam we operated the x-calibur polarimeter at the 1http://infocus.gsfc.nasa.gov/ 295 m. beilicke et al. cornell high energy synchrotron source (chess) [13]. using bragg reflection from a 2-bounce silicon (220) monochrometer a 40 kev beam was generated (2nd harmonic at 80 kev). the event rates were normalized by azimuthal acceptance on a pixel-by-pixel basis. the acceptance was determined from the response to an unpolarized beam (superposition of two perpendicular polarization planes). the energy spectrum of the compton scattered xray beam is shown in fig. 2. figure 3 shows the azimuthal scattering distribution of the polarized beam (analysis ongoing). the expected 180◦ modulation is clearly seen and the reconstructed orientation of the electric field agrees with the expected direction of the chess beam setup – confirming the functionality of the x-calibur polarimeter. detector x [m] #1 #2 #3 #4 #5 #6 #7 #8 x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x x polarized (acpt) 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 0 100 200 300 0.5 1 1.5 azimuth angle [deg]0 100 200 300 0.5 1 1.5 polarized (acpt) figure 3: x-calibur measurement of a polarized beam (energy cut: 36 − 40 kev). left: 2d map (pixel-bypixel) of counts. the (blue) boxes indicate the outline of the scintillator. right: azimuthal scattering distribution (corrected for azimuthal acceptance) for different rings. the data are fitted by a sine function. the vertical lines indicate the plane of the e field vector of the polarized beam. 4 discussion and conclusions we designed a hard x-ray polarimeter, x-calibur, and studied its performance and sensitivity. x-calibur combines a detection efficiency of close to 100% with a high modulation factor of µ ≈ 0.5, as well as a good control over systematic effects. x-calibur was successfully tested/calibrated in the laboratory and at the chess x-ray facility [13]. a 1-day x-calibur/infocus balloon flight is funded by nasa (nnx12ad51g) for fall 2013. our tentative observation program includes galactic sources (crab, her x-1, cyg x-1, grs 1915, exo 0331) and one extragalactic source (mrk 421) for which sensitive polarization measurements will be carried through. we envision follow-up longer duration balloon flights (from the southern hemisphere). successful balloon flights would motivate a satellite-borne hard x-ray polarimetry mission. acknowledgement we are grateful for nasa founding from grant nnx10aj56g & nnx12ad51g and discretionary funding from the mcdonnell center for the space sciences to build the x-calibur polarimeter. polarization measurements: this work is based upon research conducted at the cornell high energy synchrotron source (chess) which is supported by the national science foundation and the national institutes of health/national institute of general medical sciences under nsf award dmr-0936384. q.guo thanks the chinese scholarship council from china for the financial support (no.2009629064). references [1] krawczynski, h., et al.: 2011, aph. 34, 550. [2] schnittman, j.d., et al.: 2010, apj. 712, 908. doi:10.1088/0004-637x/712/2/908 [3] weisskopf, m. c., et al.: 1978, apj. 220, l117. doi:10.1086/182648 [4] dean, a. j., et al.: 2008, science. 321, 1183. doi:10.1126/science.1149056 [5] laurent, p., et al.: 2011, science. 332, 438. doi:10.1126/science.1200848 [6] http://heasarc.gsfc.nasa.gov/docs/gems. [7] tajima, h., et al.: 2010, spie 7732, 34. [8] krawczynski, h.: 2012, apj 744, 30. doi:10.1088/0004-637x/744/1/30 [9] wulf, e. a., et al.: 2007, nima 579, 371. doi:10.1016/j.nima.2007.04.085 [10] ogasaka, y, et. al.: 2005, spie 5900, 217. [11] http://geant4.cern.ch/ [12] guo, q, et. al.: 2010, arxiv 1101.0595. [13] http://www.chess.cornell.edu/ 296 http://dx.doi.org/10.1088/0004-637x/712/2/908 http://dx.doi.org/10.1086/182648 http://dx.doi.org/10.1126/science.1149056 http://dx.doi.org/10.1126/science.1200848 http://dx.doi.org/10.1088/0004-637x/744/1/30 http://dx.doi.org/10.1016/j.nima.2007.04.085 design and tests of the hard x-ray polarimeter x-calibur discussion roland walter’s comment: according to the current schedule the flight model of the hard x-ray polarimeter polar (http://www.isdc.unige.ch/polar/) should be delivered to china this year. herman marshall: how is the unpolarized flux generated? perhaps by rotating the polarization plane or by a radioactive source? matthias beilicke: we super-imposed data with the orientation of the polarization plane of 90◦ and 0◦ and independently confirmed the findings with a radiactive source (unpolarized). a 297 introduction x-calibur design performance measurements discussion and conclusions 251 acta polytechnica ctu proceedings 1(1): 251–254, 2014 251 doi: 10.14311/app.2014.01.0251 testing f(r)-theories by binary pulsars mariafelicia de laurentis1,2, ivan de martino2,3 1dipartimento di scienze fisiche, università di napoli ”federico ii” 2infn sez. di napoli compl. univ. di monte s. angelo, edificio g, via cinthia, i-80126 napoli, italy 3departamento de fisica teorica, universidad de salamanca, 37008 salamanca, spain corresponding author: felicia@na.infn.it abstract using the post-keplerian parameters to obtain, in the minkowskian limit we obtain constraints on f(r)-theories of gravity from the first time derivative of the orbital period of a sample of binary stars. in the approximation in which the theory is taylor expandable, we can estimate the parameters of an an analytic f(r)-theory, and fulfilling the gap between the general relativity prediction and the one cames from observation, we show that the theory is not ruled out. keywords: gravitation binary pulsar systems f(r)-theories gravitational waves. 1 introduction astrophysical systems like neutron stars (ns), coalescing binary systems, black holes (bhs), and white dwarfs (wds), are the most promising to study the gravitational waves (gws) emision. indeed, studying the binary system b1913+16, known as the hulsetaylor binary pulsar, the first time derivative of the orbital period was measured to be different from zero [13, 18], as predicted by general relativity (gr) when gravitational radiation is emitted. this measurements was confirmed by study in other relativistic binary systems. the agreement between gr and the observation is at the order of ∼ 1%. however, using the extended theories of gravity (etg) it should be possibile to explain the observational results as shown in [10, 12] where, starting from a class of analytic f(r)-theories it is possible evaluate the first time derivative of the orbital period and compare it with the data. this approach permit both to test the etgs both to explain the gap between observation and the theoretical prediction. this paper was organized as follow: in sec. 1 we calculate the quadrupole emission for an analytic f(r)-lagrangian using the weak-field limit; in sec 2 we compare the theoretical prediction with the observed data. finally in sec 3 we give our conclusions and remarks. 2 the first time derivative of the orbital period in the f(r)-theories the simplest extension to gr is the f(r)-gravity, in which, the lagrangian is an arbitrary function of ricci scalar [2]. starting from the field equations in f(r)gravity (for details see [2, 16, 4, 5]) f′(r)rµν − f(r) 2 gµν −f′(r);µν + +gµν�gf ′(r) = x 2 tµν , (1) 3�f′(r) + f′(r)r− 2f(r) = x 2 t , (2) where tµν = −2 √ −g δ( √ −glm) δgµν is the energy momentum tensor of matter (t is the trace), x = 16πg c4 is the coupling, f′(r) = df(r) dr , �g = ;σ ;σ, and � = ,σ ,σ. it is possibile, in the minkowskian approximation of an analytic f(r)-lagrangian1, f(r) = ∑ n fn(r0) n! (r−r0)n ' ' f0 + f′0r + f′′0 2 r2 + ... (3) to compute the quadrupole emission due to gws [10, 11]. furthermore, it is possible calculate the energy 1for convenience we will use f instead of f(r). all considerations are developed here in metric formalism. from now on we assume physical units g = c = 1. 251 http://dx.doi.org/10.14311/app.2014.01.0251 mariafelicia de laurentis, ivan de martino momentum tensor of gravitational field in f(r)-gravity that assumes the following form tλα = f ′ 0k λkα ( ḣρσḣρσ ) ︸ ︷︷ ︸ gr − 1 2 f′′0 δ λ α ( kρkσḧ ρσ )2 ︸ ︷︷ ︸ f(r) . (4) to be more precise, the first term, depending on the choice of the constant f′0, is the standard gr term, the second is the f(r) contribution. it is worth noticing that the order of derivative is increased of two degrees consistently to the fact that f(r)-gravity is of fourthorder in the metric approach [10]. in this contest, we can write the total average flux of energy due to the gws integrating over all possible directions as 〈 de dt 〉 ︸ ︷︷ ︸ (total) = g 60 〈 f′0 (... q ij... qij ) ︸ ︷︷ ︸ gr −f′′0 (.... q ij.... q ij ) ︸ ︷︷ ︸ f(r) 〉 , (5) where we point out that for f′′0 → 0 and f′0 → 4 3 , the previous equation becomes 〈 de dt 〉 ︸ ︷︷ ︸ (gr) = g 45 〈... q ij ... qij 〉 , (6) that is the prediction of gr [14, 17]. in order to evaluate the above expressions for the flux it is necessary to form explicit expressions for 〈... q ij... qij 〉 and〈.... q ij.... q ij 〉 for the system under consideration. for our purposes we consider a binary pulsar system. if we assume a keplerian motion of the stars in the binary system, wherewe mp is the pulsar mass, mc the companion mass, and µ = mcmp mc + mp is the reduced mass, it is possible to compute the time average of the radiated power computing the first time derivative of the orbital period [11] ṗb = − 3 20 ( t 2π )−5 3 µg 5 3 (mc + mp) 2 3 c5(1 − �2) 7 2 × × [ f′0 ( 37�4 + 292�2 + 96 ) − f′′0 π 2t−1 2(1 + �2)3 × × ( 891�8 + 28016�6 + 43520�2 + 3072 )] , (7) where � is the orbital eccentricity and t is the orbital period of the binary. 3 methodology and data analysis knowing exactly the lagrangian that describes the system, we can predict the orbital period decay, however,we want understand how well the relativistic binary systems can fix bounds on f(r) parameters using eq. (7), and getting an estimation of the second derivative of the lagrangian with respect to ricci scalar, f′′0 . we use the following prescription, the difference between the first derivative of the binary observed period variation (ṫbobs ±δ) and the theorethical one obtained by gr, ∆ṫb = ṫbobs − ṫgr, is fulfilled imposing that: ṫbobs − ṫgr −f ′′ 0 ṫbf(r) = 0, (8) ṫbobs ±δ − ṫgr −f ′′ 0±δ ṫbf(r) = 0, (9) where δ is the experimental error, that we propagate on the ṫbobs, into an uncertainty on f ′′ 0±δ . in this way, the extra contribution to the loss of energy due to the emission of gws radiation in the etgs regime can provide to fill the difference between theory and observations. we select a sample of observed relativistic binary pulsars (see their references reported in tab. 1 of [11]) for which we compute the correction ṫbf(r) , the difference ∆ṫgr between ṫbobs and ṫgr (equal to the correction −f′′0 ṫbf(r) ), the corresponding f ′′ 0 solution of (8), the interval centered on f′′0 and finally, the interval centered on f′′0 and computed from the difference: f′′0+δ −f′′0−δ 2 , all results are reported in tab. 1. in fig. 1 we show, for sake of convenience, in logarithmic scale, the absolute values of f′′0 reported in tab.1 versus the ratio ṫbobs ṫgr . there are six binaries in tables, for which the etgs are not ruled out 0.04 ≤ f′′0 ≤ 38, getting 0.5 ≤ ṫbobs ṫgr ≤ 1.5. for those systems the difference between ṫgr and ṫbobs can be explained adding an extra contribution that comes out from the f(r)-thoery. instead for most of binaries we have f′′0 values that can surely rule out the theory, since taking account of the weak field assumption we obtain 38 ≤ f′′0 ≤ 4 × 107. from this last values to the first ones, there is a jump of about four up to five order of magnitude on f′′0 . the origin of these strong discrepancies, perhaps, is due to the extreme assumption we made, to justify the difference between the observed ṫbobs and the predicted ṫgr using the etgs. 252 testing f(r)-theories by binary pulsars table 1: upper limits of f′′0 correction to ṫgr of binary relativistic pulsars assuming that all the loss of energy is caused by gravitational wave emission. we reported the j-name of the system,the difference ∆ṫgr between ṫbobs and ṫgr equal to the correction −f ′′ 0 ṫbf(r) , the correction ṫbf(r) , the corresponding f ′′ 0 solution of (8), the interval centered on f′′0 and computed from the difference f′′0+δ −f′′0−δ 2 ,where f′′0±δ ,are the corresponding solutions of ( 8) taking account of the experimental errors ±δ on the observed orbital period variation ṫbobs. name ∆ṫgr ṫbf(r) f ′′ 0 ±∆f′′0 j2129+1210c -2.17e-14 6.01e-13 3.61e-02 8.32e-02 j1915+1606 -2.04e-14 2.10e-13 9.74e-02 4.77e-03 j0737-3039a -4.23e-15 1.86e-14 2.28e-01 9.15e-02 j1141-6545 -1.65e-14 3.88e-15 4.25e+00 6.44e+00 j1537+1155 5.39e-14 1.42e-15 -3.79e+01 7.03e-02 j1738+0333 -1.56e-15 1.06e-16 -1.47e+01 2.92e+01 j0751+1807 1.41e-13 8.98e-16 -15.7e+01 1.002e+01 j0024-7204j -5.22e-13 3.13e-16 1.67e+03 4.15e+02 j1701-3006b -5.03e-12 8.81e-16 5.71e+03 7.04e+01 j2051-0827 -1.55e-11 4.77e-16 3.24e+04 1.68e+03 j1909-3744 -5.47e-13 2.62e-18 2.09e+05 1.14e+04 j1518+4904 2.41e-13 3.42e-19 -7.05e+05 6.43e+03 j1959+2048 1.47e-11 1.07e-17 -1.38e+06 7.51e+04 j2145-0750 4.01e-13 1.00e-19 -4.00e+06 2.99e+06 j0437-4715 1.59e-13 1.04e-19 -1.57e+06 2.73e+06 j0045-7319 3.02e-07 1.11e-16 2.74e+9 8.13e+07 j2019+2425 -3.00e-11 1.11e-22 2.71e+11 5.41e+11 j1623-2631 4.00e-10 2.02e-23 -1.98e+13 2.97e+13 4 discussion and remarks in this paper, we develop expressions for quadrupole gravitational radiation in f(r)-gravity theory using the weak field technique and apply these results, which are applicable in general, to a sample of a binary pulsars, though their orbits are eccentric. here, we seen that, where the gr theory is not enough to explain the gap between the data and the theoretical estimation of the orbital decay, there is the possibility to extend the gr theory with a generic f(r)theory to cover the gap. according to eq. (7),we have selected a sample of relativistic binary systems for which the first derivative of the orbital period is observed, we have computed the theoretical quadrupole radiation rate, and finally we have compared it to binary system observations. from tab. 1, it is seen that the first five systems have masses determined in a manner quite reliable, while for the remaining sample, masses are estimated by requiring that the mass of the pulsar is 1.4m� and, assuming for the orbital inclination one of the usual statistical values (i = 60◦ or i = 90◦ ), and from here comes then the estimate of the mass of the companion star. so a primary cause of major discrepancies, not only for the etgs, but also for the gr theory, between the variation of the observed orbital period and the predicted effect of emission of gravitational waves, could be a mistake in the estimation of the masses of the system. in addition, other causes may be attributable to the evolutionary state of the system, which, for instance, if it does not consist of two neutron stars may transfer mass from companion to the neutron star. in our sample, there are only five double ns that can be used to test gr and etgs. taking into account of the strong hypothesis we made, the etg correction to ṫgr can also include the galactic acceleration term correction ([7], [8]). here, we give a preliminary result about the energy loss from binary systems and we show that, when the nature of the binary systems can exclude energy losses due to trade or loss of matter, then, we can explain the gap between the first time derivative of the observed orbital period and the theoretical one predicted by gr, using an analytical f(r)-theory of gravity. 253 mariafelicia de laurentis, ivan de martino 10 0 10 5 10 10 10 15 10 −4 10 −2 10 0 10 2 10 4 10 6 10 8 10 10 f′′ 0 ṫ b o b s ṫ g r ṫbobs ṫgr = 1.5 ṫbobs ṫgr = 0.5 f ′′ 0 = +0.04 f ′′0 = 38 psrj1537+1155 psrj1738+0333 f ′′ 0 = 14.71 psrj2129+1210c figure 1: in figure there are shown, for sake of convenience, in logaritmic scale, the absolute values of f′′0 reported in tab. 1 versus the ratio ṫbobs ṫgr . we must note that for five binaries the etgs we are probing is not ruled out 0.04 ≤ f′′0 ≤≈ 38, for those systems the difference between ṫgr and ṫbobs is tiny, indeed we get 0.5 ≤ ṫbobs ṫgr ≤ 1.5. instead for most of binaries we have f′′0 values that can surely rule out the theory, since taking account of the weak field assumption we obtain 38 ≤ f′′0 ≤ 4 × 107. from this last values to the first ones, there is a jump of about four up to five order of magnitude on f′′0 . references [1] bogdanos c., capozziello s., de laurentis m., nesseris s., 2010, astrop. phys. , 34, 236. doi:10.1016/j.astropartphys.2010.08.001 [2] capozziello s., de laurentis m., 2011, physics reports 509, 167. doi:10.1016/j.physrep.2011.09.003 [3] capozziello s., corda c., de laurentis m., 2008 phys. lett. b 669, 255-259 . doi:10.1016/j.physletb.2008.10.001 [4] capozziello s., francaviglia m., 2008, gen. rel. grav. 40,357. doi:10.1007/s10714-007-0551-y [5] capozziello s., de laurentis m., faraoni v., 2009 ,the open astr. jour , 2, 1874. 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[18] weisberg j.m., nice d.j. and taylor j.h., 2010, apj, 722,1030-1034. doi:10.1088/0004-637x/722/2/1030 [19] will c.m, 1993, ”theory and experiment in gravitational physics” cambridge university press, uk. doi:10.1017/cbo9780511564246 254 http://dx.doi.org/10.1016/j.astropartphys.2010.08.001 http://dx.doi.org/10.1016/j.physrep.2011.09.003 http://dx.doi.org/10.1016/j.physletb.2008.10.001 http://dx.doi.org/10.1007/s10714-007-0551-y http://dx.doi.org/10.1103/physrevd.58.042001 http://dx.doi.org/10.1016/j.astropartphys.2011.08.006 http://dx.doi.org/10.1093/mnras/stt216 http://dx.doi.org/10.1111/j.1365-2966.2012.21253.x http://dx.doi.org/10.1093/acprof:oso/9780198570745.001.0001 http://dx.doi.org/10.1016/j.physrep.2011.04.001 http://dx.doi.org/10.1088/0004-637x/722/2/1030 http://dx.doi.org/10.1017/cbo9780511564246 introduction the first time derivative of the orbital period in the f(r)-theories methodology and data analysis discussion and remarks 132 acta polytechnica ctu proceedings 1(1): 132–138, 2014 132 doi: 10.14311/app.2014.01.0132 auger highlights antonella castellina 1 for the pierre auger collaboration 2 1istituto nazionale di astrofisica oato and infn torino 2observatorio pierre auger, av. san mart́ın norte 304, 5613 malargüe, argentina (full author list : http://www.auger.org/archive/authors 2012 12.html) corresponding author: castellina@to.infn.it abstract the pierre auger observatory has been designed to investigate the origin and nature of the ultra high energy cosmic rays using a hybrid detection technique. a review of selected results is presented, with the emphasis given to the measurement of energy spectrum, mass composition and arrival directions. keywords: ultra high energy cosmic rays sources energy spectrum primary composition. 1 introduction understanding the sources, nature and propagation properties of the cosmic rays at ultra high energies (e > 1018 ev) is one of the key questions in astroparticle physics. from the experimental point of view, their study can be performed indirectly, by exploiting the extensive air showers they produce by interacting with the nuclei in the earth atmosphere. among the different features characterizing the spectral shape, the region between ' 1018−1019 ev is thought to host the transition from galactic to extragalactic cosmic rays. different models explain it as due to e+/e− pair production of protons with the photons of the cosmic microwave background (cmb) [1], or more traditionally to the intersection of a steep galactic component and the onset of a flatter extragalactic one [2]. at even higher energies, above ' 4 1019 ev, a cut-off in the cosmic ray flux is expected, due to photo-pion production of extragalactic protons in the cmb (the ”gzk cut-off” [3]) , although the same feature could arise when reaching the limits in the maximum energy of the sources. however, the all particle spectrum cannot provide a discrimination among the different hypotheses, and the determination of the primary composition is mandatory to reach any reliable conclusion. the analysis of the arrival directions and their anisotropy can give further insight into the sources and provide information about the magnetic fields which the ultra high energy (uhe) cosmic rays experience during their travel to earth. the pierre auger observatory has been specifically designed to investigate the origin and the nature of ultra high energy cosmic rays. it is located in malargüe, argentina, and consists of a surface array (sd) of 1660 water cherenkov stations on an area of ' 3000 km2, overlooked by 27 air fluorescence telescopes (fd) grouped in four sites [4]. thanks to the possibility of combining the information from the surface array, measuring the lateral distributions of secondary particles at the ground, and the fluorescence telescopes, observing the longitudinal profile, the reconstruction capabilities are enhanced with respect to the individual detector components. 2 the energy spectrum the energy spectrum above 2.5 1018 ev has been determined using the data from the sd [5], considering only events up to 60◦. the exposure is obtained by integrating the number of active stations over time; the overall acceptance uncertainty is ' 3% [6]. the energy calibration is derived directly from data, using a subset of high quality hybrid events, i.e. events reconstructed by both the fd and the sd [7]. despite the low duty cycle of the fd, the energy spectrum could be extended to 1018 ev using hybrid events, thanks to the good energy resolution and low threshold, thus investigating the transition region in detail [8]. the total systematic uncertainty in the energy scale is about 22%, the main contribution coming from the uncertainty in the fluorescence yield (14%) and in the reconstruction of the longitudinal profile (10%). the sd and hybrid spectra can be combined using a maximum likelihood method, since both have the same systematic uncertainties in the energy scale. the normalization uncertainties are on the contrary independent and have been used as additional constraints in the 132 http://dx.doi.org/10.14311/app.2014.01.0132 auger highlights procedure. the resulting spectrum is shown in fig.1; a fit with three power laws is shown by the dashed lines, while the solid line indicates the result of a fit with two power laws and a smooth function. the ankle feature is present at an energy of 1018.62 ev; the cutoff is clearly seen with a significance of 20 σ. different astrophysical models can be compared to our data; however, the energy spectrum can be described by both a heavy or proton composition at the highest energies and the information must be complemented by independent measurements of the primary composition. figure 1: the combined auger energy spectrum. only statistical uncertainties are shown. the systematic uncertainty on the energy scale is 22%. a comparison of the auger results with data from hires, telescope array and yakutsk has been recently performed [9]. the various fluxes can be rescaled assuming that any difference among them be due solely to energy scale and not to aperture calculations or energy resolution. the differences found are entirely consistent with the systematic energy uncertainties quoted by the experiments. 3 the nature of the primaries the most direct information about composition can be obtained by measuring the longitudinal development of showers in hybrid events, thus determining the mean depth of maximum development, xmax, and its fluctuation, rms(xmax). for each event, xmax depends on the depth of first interaction of the primary in the atmosphere and on the subsequent development of the shower; for this reason, the interpretation of the results in terms of composition is complicated by the uncertainties in the hadronic interaction models used in simulations. about six years of hybrid data have been analyzed applying fd quality cuts and ensuring that no bias with respect to the cosmic ray composition is introduced in the data sample [10]. having been corrected for the detector resolution, the xmax and its fluctuations are detector independent and can be directly compared to the predictions of different models, as shown in fig.2. both observables show a change for e > 5 1018 ev towards an increasingly heavy composition in comparison to the model predictions. the average resolution of xmax ' 20 g/cm2 in the considered energy range. figure 2: < xmax > and rms(xmax) as a function of energy, compared with the predictions of air shower simulations using different hadronic interaction models. in general, different values for rms(xmax) are allowed for different combinations of elements [11]. the fluctuations predicted by the considered hadronic interaction models and shown in fig.2 are evaluated only for pure compositions . a different conclusion, leading to a light composition up to the highest energies, has been drawn from the data of the hires and telescope array. however, a direct comparison of their results with the auger ones is not possible, as the detector biases are included in their simulation. furthermore, their dataset is smaller that that of auger. a lenghty discussion about this comparison can be found in [12] (and refs. therein). starting from an extension of the heitler model of ex133 antonella castellina tensive air showers [13], a method for interpreting the results of xmax and rms(xmax) in terms of mass composition has been developed. as discussed in [14], xmax is only function of the mean logarithmic mass < ln a >, and as such it carries information on the average composition. on the contrary, both the shower-to-shower fluctuations and the dispersion in the mass distribution contribute to rms(xmax). this information can be used to extract the < lna > and its variance from the observables and to build the plots shown in fig.3, where the size of the data points increases with increasing energy. the energy evolution of the composition is common to all models; reduced systematic uncertainties will allow in the future to test or even exclude some of them. different consequences from the astrophysical point of view can be derived from this comparison. extragalactic sources of protons seem to be disfavored by our composition result, within the uncertainties on the hadronic interaction models used to interpret the data. in a propagation scenario, nuclei from nearby sources could produce small mass dispersion at earth, as propagation would not be able to degrade mass and energy. if on the other hand the proton component is depleted by the reach of a rigidity dependent end of the injection spectrum, and if sources are uniformly distributed, hard injection spectra with low energy cutoff, together with local sources, could explain the data [15, 16]. figure 3: the pierre auger data in the (< lna > ,σ2lna) plane for different hadronic interaction models. grey contours limit the allowed region, the systematic uncertainties are shown by the black lines. 4 primary photons and neutrinos uhe primary photons and neutrinos can provide invaluable information about the astrophysics of cosmic rays. their detection would be a direct proof of the gzk cutoff; limits on exotic models [17] and tests for new physics [18] could be obtained from a positive or negative result on their detection. in both cases, their search is based on the characteristic features of the showers they produce in comparison to the hadronic ones. primary photons produce late developing showers, a characteristic further enhanced by the lpm effect [19]. the deeper xmax, observable by the fd, is associated to a more dispersed distribution of the arrival time of the particles at ground level. at a given distance from the shower axis, the arrival time of the first particles is delayed with respect to a planar shower front and the radius of curvature is thus expected to decrease for photon induced showers. these observables can be recorded by means of the sd. the upper limits derived from both the sd and the hybrid data collected by auger are shown in fig.4 and discussed in [20, 21]. astrophysical scenarios are favoured with respect to top-down models. figure 4: upper limits on the integral photon flux. different expectations are shown for comparison. primary neutrinos can produce showers characterized by a significant electromagnetic component; the huge hadronic background can be eliminated by looking at large zenith angles. ντ can interact by charged current in the earth crust, producing earth-skimming upward-going showers; neutrinos of any flavour can interact in the atmosphere by neutral or charged current giving rise to downward-going events. assuming a diffuse neutrino flux ' k e−2, 90% cl limits on their flux are obtained, as shown in fig.5. neutrinos from point sources were also searched for, over a broad declination range (north of ' −65◦ and 134 auger highlights south of ' 55◦), for eν > 1017 ev. for a differential neutrino flux ' kps e−2, 90% cl limits of ' 5 · 10−7 and 2.5 · 10−6 gev cm−2 s−1 have been obtained for up-going and down-going events respectively [22]. figure 5: differential and integrated upper limits on the single flavour e−2 ν flux (thin black lines: downward-going ν, thick red line: earth-skimming ν). different models are shown for comparison and discussed in [23] 5 anisotropies the spatial distribution of the arrival directions of uhe cosmic rays as a function of energy is a key observable to provide information about their sources and nature, complementary to those of energy spectrum and composition. particles of uhe are most probably extragalactic, and if the observed cutoff in the spectrum can be attributed to the gzk propagation effect we could expect their sources to be confined in our courtyard, within about 100 mpc. in 2007 [24] the auger collaboration reported the observation of a correlation between the arrival directions of the highest energy cosmic rays and the positions of nearby agn from the véron-cetty-véron catalogue [25]. the result came from an analysis of indipendent data with a priori parameters determined from an exploratory scan; this allowed to avoid the use of penalty factors which would be needed in a posteriori analyses. the most recent update of this search is shown in fig.6 [26]: the fraction of correlating cosmic rays is (33±5)% (28 events correlating out of a total of 84). the probability of this correlation to occur by chance if the true distribution of arrival directions is isotropic stays below 1%. the independent averages of 10 consecutive events are also shown (black dots). a recent comparison of our result with the telescope array and the yakutsk ones showed that the correlating fractions are compatible [27]. more data are necessary to show whether this correlation is statistically significant or not. figure 6: the correlating fraction as a function of the total number of time-ordered events. different confidence levels are shown, together with the isotropy value p=0.21 and the current estimate of all data, pdata = 0.33 ± 0.05. another possible scenario is that the anisotropy is dominated by cosmic rays originating from the vicinity of centaurus a, the nearest active galaxy with an estimated distance of about 3.8 mpc, since 19 events out of 7.6 expected have arrival directions within 24◦ of cena. a kolmogorov-smirnov test shows that the chance probability for this to happen is 4%. directionally aligned events, or ”multiplets”, can be expected from the same source after deflection in the magnetic fields, showing a correlation between the arrival direction and the inverse of energy. the largest multiplet found was one 12-plet, but also in this case the probability for it to come from an isotropic distribution is ' 6% [28]. potential sources of galactic cosmic rays have been looked for by performing a blind search for neutron primaries [29]. in fact, due to the relativistic time dilatation the uhe neutron mean decay length is (9.2 × e/eev ) kpc; above 2 eev, neutron emitters can be searched for in the whole galaxy. auger can detect neutron showers by a simple search for an excess of proton-like showers from a specific direction in the sky. no candidates have been found, bringing to a median flux upper limit of 0.0114 n km−2 yr−1 above 1 eev. the absence of a neutron flux from the galaxy, which could be expected in the hypothesis of sources steadily emitting protons and neutrons with similar luminosity, could be a hint that the sources at eev energy could be e.g. extragalactic, transient or weak but densely distributed. the large scale distribution of the arrival directions of cosmic rays is another fundamental tool in the search for their origin. the results from a study performed using data from the sd array are shown in fig.7 [30]. no significant anisotropies are observed, resulting in the most stringent bounds on the first harmonic amplitude above 2.5 1017 ev. 135 antonella castellina figure 7: equatorial dipole component (top) and phase of the first harmonic (bottom) as a function of energy. the obtained limits already exclude some of the galactic models (labeled in fig.7 with a and s, indicating antisymmetric and symmetric galactic magnetic fields), according to which the cosmic rays at these energies are galactic and can escape by diffusion and drift motion. in the model labeled gal, cosmic rays are assumed to be galactic at all energies, and the anisotropy is due to purely diffusive motion caused by the turbulent component of the galacic magnetic field. in extragalactic models, the transition is put at the second knee and the cosmic rays large scale distribution is influenced by the relative motion of the observer with respect to the frame of the source. assuming that the frame in which the cosmic ray distribution at these energies is isotropic is coincident with the cosmic microwave background rest frame, a small anisotropy (the extragalactic compton-getting effect, labeled c-g xgal) is expected. interestingly, the phase of the first harmonic shows a smooth transition between a common phase of ' 270◦ below 1 eev and ' 100◦ above 5 eev. a consistency of the phase in ordered energy intervals can indeed be expected in presence of a real underlying anisotropy, standing out of the background more prominently than the amplitude. however, no confidence level can for the moment be assigned to this result, being an ”a posteriori” observation. the study of the large scale anisotropy has been performed for the first time with auger data as a function of both the right ascension and the declination and expressed in terms of dipole and quadrupole amplitudes [31]. no significant deviations from isotropy are detected. under the hypothesis that any anisotropy is dominated by these moments, the 99% cl upper limits can be derived, as shown in fig.8. figure 8: 99% cl upper limits on the dipole and quadrupole momenta as a function of energy. as an example of the power of the measurement to discriminate among different astrophysical models, the experimental limits are compared in the figure with the expectations from a toy model, in which the sources of protons and iron are stationary and uniformly distributed in the galactic disk. being the expected amplitudes for protons largely above the allowed upper limits, we can exclude this scenario for the light component of eev primary cosmic rays. 6 future developments the pierre auger observatory has reached a cumulative exposure of more than 26000 km2sr yr. new information about the characteristics of the primary cosmic rays have been derived, opening at the same time more questions and pointing to the need of an 136 auger highlights extension of the life time of auger. the measurement of composition of the primary particles from the eev region up to the highest energies has emerged as the key for hitting the hottest scientific questions: a) understand the origin of flux suppression discussed in sect.2, if due to the reach of the maximum energy at injection or to the gzk effect (a clear signature of which would be the observation of a flux of primary photons and neutrinos); b) perform composition driven anisotropy searches; c) determine the energy at which the transition from galactic to extragalactic sources of cosmic rays takes place. acknowledgement the author wishes to thank the organizers for the warm hospitality in mondello and the stimulating and interesting discussions during the workshop. references [1] v. berezinski, a. gazizov, s. grigorieva: 2006, phys.rev.d 74, 043005. doi:10.1103/physrevd.74.043005 [2] a.m. hillas: 2006, in cosmology, galaxy formation and astroparticle physics on the pathway to the ska klöckner, h.-r., jarvis, m., rawlings, s. (eds.) april 10th-12th 2006, oxford, united kingdom. [3] k. greisen: 1966, phys. rev. lett. 16, 748; g.t. zatsepin and v.a. kuzmin: 1966, sov. phys. jetp lett. 4, 78. [4] the pierre auger collaboration: 2010, nucl.instr.meth. a 620, 227. doi:10.1016/j.nima.2010.04.023 [5] the pierre auger collaboration: 2008, phys.rev.lett. 101, 061101. doi:10.1103/physrevlett.101.061101 [6] the pierre auger collaboration: 2010, nucl.instr.methods a613, 29. [7] r. pesce for the pierre auger collaboration: 2011, in 32nd int.cosmic ray conf., beijing, china; arxiv:1107.4809. [8] m. settimo for the pierre auger collaboration: 2012, eur.phys.j. plus 127, 87. [9] b. dawson et al. for the pierre auger, telescope array and yakutsk collaborations: 2012, in int.symposium on future directions in uhecr physics, cern. [10] p. facal for the pierre auger collaboration: 2011, in 32nd int.cosmic ray conf., beijing, china; arxiv:1107.4804. [11] k.-h. kampert, m. unger: 2012, astrop.phys. 35, 660. [12] e. barcikowski et al. for the pierre auger, telescope array and yakutsk collaborations: 2012, in int.symposium on future directions in uhecr physics, cern. [13] j. matthews: 2005, astrop.phys. 22, 387. [14] the pierre auger collaboration: 2013, jcap 02, 026. [15] a.m. taylor, m. ahlers, f.a. aharonian: 2011, phys.revd84, 105007. doi:10.1103/physrevd.84.105007 [16] d. allard: 2012, astrop.phys. 39-40, 33. doi:10.1016/j.astropartphys.2011.10.011 [17] p. battacharije and g. sigl: 2000, phys.rep.327, 109. [18] m. galaverni and g. sigl: 2008, phys.rev.lett.100, 021102. [19] l.d. landau, i.ya. pomeranchuk: 1953, dokl.acad.nauk.92, 553; a.b. migdal: 1956, phys.rev.103, 1811. [20] the pierre auger collaboration: 2009, astrop. phys.31, 399; 2008, astrop. phys.29, 243. [21] v. scherini for the pierre auger collaboration: 2012, in int.symposium on future directions in uhecr physics, cern. [22] the pierre auger collaboration: 2012, apj.lett. l4, 755. [23] the pierre auger collaboration: 2013, adv.in high energy phys.res., 2013, 708680. [24] the pierre auger collaboration: 2007, science 318, 938. [25] m.p. véron-cetty, p. véron: 2006, a&a 445, 773. [26] k.-h. kampert for the pierre auger collaboration: 2011, highlight talk in 32nd int. cosmic ray conf., beijing, china. [27] o. deligny et al. for the pierre auger, telescope array and yakutsk collaborations: 2012, in int.symposium on future directions in uhecr physics, cern. 137 http://dx.doi.org/10.1103/physrevd.74.043005 http://dx.doi.org/10.1016/j.nima.2010.04.023 http://dx.doi.org/10.1103/physrevlett.101.061101 http://dx.doi.org/10.1103/physrevd.84.105007 http://dx.doi.org/10.1016/j.astropartphys.2011.10.011 antonella castellina [28] the pierre auger collaboration: 2012, astrop.phys. 35, 354. doi:10.1016/j.astropartphys.2011.10.004 [29] the pierre auger collaboration: 2012, apj 760, 148. doi:10.1088/0004-637x/760/2/148 [30] the pierre auger collaboration: 2011, astrop.phys. 34, 627. doi:10.1016/j.astropartphys.2010.12.007 [31] the pierre auger collaboration: 2012, apj.suppl. 203, 34; 2013, apj.lett. 762, l13. 138 http://dx.doi.org/10.1016/j.astropartphys.2011.10.004 http://dx.doi.org/10.1088/0004-637x/760/2/148 http://dx.doi.org/10.1016/j.astropartphys.2010.12.007 introduction the energy spectrum the nature of the primaries primary photons and neutrinos anisotropies future developments 94 acta polytechnica ctu proceedings 2(1): 94–98, 2015 94 doi: 10.14311/app.2015.02.0094 supersoft x-ray source cal 83: a possible ae aqr-like system a. odendaal1, p. j. meintjes1, p. a. charles2, a. f. rajoelimanana3 1department of physics, university of the free state, p.o. box 339, bloemfontein, 9300, south africa 2school of physics and astronomy, southampton university, southampton so17 1bj 3department of physics, north-west university, private bag x2046, mmabatho 2735, south africa corresponding author: winka@ufs.ac.za abstract cal 83 is a close binary supersoft x-ray source in the large magellanic cloud. a ∼67 s periodicity detected in supersoft x-rays is most probably associated with the spin period of a highly spun-up white dwarf (wd). the variability in the period is ascribed to the obscuration of the wd by the hydrogen burning envelope surrounding it, rotating with a period that is close to, but not quite synchronized with, the wd rotation period. optical spectra obtained with salt exhibit accretion disc emission lines with broad wing structures and p cyg profiles, indicating mass outflows. timing analysis of photometrical observations performed at the south african astronomical observatory (saao) revealed variable signals at ≤1 mhz which are thought to be associated with quasi-periodic oscillations from an accretion disc. the short spin period inferred for cal 83 can be the result of spin-up by accretion disc torques during a long mass transfer history, placing this source on a similar evolutionary track as the cataclysmic variable ae aqr. keywords: stars: individual: cal 83 binaries: close white dwarfs stars: oscillations stars: winds, outflows stars: evolution. 1 introduction supersoft x-ray sources (sss) are highly luminous in the supersoft x-ray band, with >90% of the unabsorbed x-ray flux being below ∼0.5 kev. these sources were first discovered by the einstein x-ray observatory and rosat (see kahabka & van den heuvel (2006) and references therein). van den heuvel et al. (1992) (hereafter vdh92) showed that the low effective temperatures and high luminosities of sss can be explained by the nuclear burning of hydrogen on the surface of a white dwarf (wd) accreting from a binary companion (the close binary sss model). the accretion rate required for steady nuclear burning is ∼ 1 − 4 × 10−7 m� yr−1. such a high accretion rate can be sustained if the companion mass is comparable to or greater than the wd mass, as this will cause the roche lobe of the donor to shrink and drive mass transfer on the thermal time-scale of the donor. cal 83 is a close binary sss in the large magellanic cloud. however, it is not a persistent x-ray source and several x-ray off-states have been observed, with long-term optical variability anti-correlated with the long-term x-ray variability, as described by e.g. rajoelimanana et al. (2013). these authors also derived a refined orbital period of 1.047529(1) d for cal 83 from ogle-iii photometry. the inclination has been estimated to be in the i = 20 − 30◦ range (see cowley et al. (1998) and references therein, hereafter co98). lanz et al. (2005) used nlte models during a combined analysis of xmm-newton and chandra data, and their results indicate a massive wd (mwd ∼ 1.3 ± 0.3 m�). schmidtke & cowley (2006) found a periodic signal of 38.4-minutes in the chandra letg data of cal 83, which they ascribed to possible non-radial pulsations in the accreting wd. the optical spectrum of cal 83 is characterized by balmer and he ii accretion disc emission lines (co98 and references therein). the he ii λ4686 emission line has a broad, variable wing structure that has also been observed in some of the balmer lines. in each particular spectrum, these wing structures are either all towards the blue side of the main components of these emission lines, or all towards the red side. according to crampton et al. (1987) (hereafter cr87), this may be associated with the slow precession of an accretion disc with outflows through either wind or a weakly collimated jet. the ∼67 s x-ray periodicity is briefly summarized in §2.1 (a detailed discussion is provided in odendaal et al. 2014). preliminary results of the analysis of optical data are presented in §2.2 and §2.3. after this discussion, the possible evolutionary scenario of cal 83 is brought into context with that of the cv ae aqr in §3, followed by the conclusions in §4. 94 http://dx.doi.org/10.14311/app.2015.02.0094 supersoft x-ray source cal 83: a possible ae aqr-like system 2 cal 83: observations and results 2.1 the ∼67 s x-ray pulsation the xmm-newton archive contains 23 observations of cal 83, 4 of which were during an x-ray off-state. a systematic search for short time-scale periodicities that may be associated with the wd spin was performed on the 19 on-state light curves. the starlink period1 package was used to obtain a lomb-scargle (ls) periodogram of each light curve. a power peak at a period of 66.8 ± 0.4 s was discovered at a > 99.9999% significance level in the periodogram of the epic pn light curve of observation 0506531501, and at a lower significance level in 6 of the other pn periodograms. in 3 of these observations, 1 of the mos light curves also exhibited a weak ∼67 s periodicity. the ls periodogram of observation 0506531501, as well as the light curve folded on the 66.8 s period, are shown in fig. 1. 0 10 20 30 40 0 15 30 45 60 75 90 l o m b -s c a rg le p o w e r frequency (mhz) 99.73% -0.60 +0.60 7.98 0.0 0.5 1.0 1.5 2.0 c o u n ts s -1 phase figure 1: lomb-scargle periodogram (left) and folded light curve (right) of cal 83 xmm-newton dataset 0506531501 pn. both the periodogram and folded light curve display a significant 66.8 s oscillation. the overall significance of this detection, considering the pn datasets of all 19 the on-state observations, is ≥3σ. the detected period exhibits a spread of ±3 s around ∼67 s between different observations, and even within a single observation. this variability cannot be explained by doppler shifts due to wd orbital motion. the long-term presence of the pulsation suggests its association with the wd spin, although the variation in the detected period complicates matters. it is possible, however, that we are observing the wd spin, but through an extended envelope around the accreting wd. according to ibragimov et al. (2003), the photospheric radius in a steady nuclear burning sss may have a radius 2-3 times the radius of the wd itself. the spread in the observed period can possibly be explained by the envelope not being quite synchronized with the wd, resulting in a slippage effect and associated variation in the pulsation period. 2.2 optical spectroscopy with salt a series of optical spectra of cal 83 has been obtained with the robert stobie spectrograph (rss) on the southern african large telescope (salt), using the pg0900 grating to obtain a resolution of r ∼ 1500. these spectra show similar characteristics to those discussed in §1, with strong balmer and he ii emission lines. the velocity widths of these lines indicate emission from regions compatible with an accretion disc around the wd. these lines also exhibit blue wing structures extending to ∼2000 km s−1 from the line centres, supporting the existence of a wind or weakly collimated jet in the system. p cyg profiles are also evident in the hα and hβ lines, serving as further evidence of mass outflow. shown in fig. 2 are the hα and hβ profiles from a spectrum obtained on 2 february 2013. 1000 1250 1500 1750 2000 −5000 −2500 0 2500 c o u n ts velocity (km s −1 ) hβ λ4861 1000 2000 3000 4000 −5000 −2500 0 2500 velocity (km s −1 ) hα λ6563 figure 2: hβ and hα profiles of cal 83, obtained with salt on 2 feb 2013. for mwd ∼ 1.3 m�, the wd radius is rwd ≈ 4 × 108(mwd/1.3 m�)−0.8 cm (see eracleous & horne 1996 and references therein). the mass function can be determined if the semi-amplitude (k1) of a spectral line originating close to the wd is known. although the orbital coverage of the salt spectra was not adequate to calculate a new k1, various values have been reported in the literature. adopting k1 ∼ 35 km s−1, i ∼ 25◦ (see co98) and porb = 1.047529 d, the mass function is 4.7×10−3 m�, and a secondary mass of m2 ∼ 0.61 m� is obtained, yielding a mass ratio q = m2/mwd ∼ 0.47. however, this value of k1 was derived from the radial velocity modulation of the he ii λ4686 disc emission line, and this line would only represent the wd orbital motion if the he ii emission was distributed symmetrically about the wd. as discussed by e.g. kaitchuck et al. (1994), there are some problems with this assumption, compromising the mass function calculation. 2.3 fast photometry with shoc in april 2013, fast photometrical observations were obtained with one of the sutherland high-speed optical 1www.starlink.rl.ac.uk/star/docs/sun167.htx/sun167.html 95 a. odendaal et al. cameras (shoc2) on the saao 1.9-m telescope at sutherland. these cameras have been commissioned recently, and are optimized for high time resolution photometry. a clear filter was used for the cal 83 observations, and with this sensitive instrument, exposure times in the 0.5-5 s range were adequate, depending on observing conditions. aperture photometry was utilized to obtain light curves, and the cal 83 light curves were corrected by performing differential photometry using 4 comparison stars. the starlink period package was used to perform a ls analysis. the ls periodograms of cal 83 did not reveal any significant pulsation at the inferred spin period of ∼67 s or at higher frequencies. however, some of the cal 83 periodograms exhibited a significant peak in the region just below 1 mhz. these peaks have broadened profiles, and their positions do not stay constant. in fig. 3, the ls periodogram of the shoc observation on 15 april 2013 is presented, showing a peak at 1168±214 s. the light curve folded on this period is also shown. this observation had a total length of 3245 s, and the light curve was binned to 20 s for this analysis. these peaks are ascribed to quasi-periodic oscillations associated with the accretion disc around the wd. 0 3 6 9 12 15 0 2 4 6 8 10 l o m b -s c a rg le p o w e r frequency (mhz) 99.73% -7.650 -7.625 -7.600 -7.575 -7.550 -7.525 0.0 0.5 1.0 1.5 2.0 d if fe re n ti a l m a g n it u d e phase figure 3: shoc periodogram and light curve folded on the detected period of 1168 ± 214 s (obtained on 15 april 2013). these quasi-periods at ≤1 mhz may be the keplerian periods of blobs of material orbiting the wd in the accretion disc, or possibly beat periods between the presumed spin period of p∗ ∼ 67 s and blobs orbiting in the inner regions of the disc. assuming that the 1168±214 s periodicity is a keplerian period, the associated keplerian radius is 46 rwd. on the other hand, it may be a beat period between p∗ ∼ 67 s and a keplerian period of ∼71 s associated with blobs orbiting at ∼7 rwd. adopting the orbital parameters in §2.2 yields a circularization radius of ∼150 rwd, supporting the compatibility of both the keplerian radii above with emission from the accretion disc. 3 the evolution of cal 83: an ae aqr analogue? ae aqr is a nova-like variable with a short wd spin period (33.08 s), and a long orbital period (9.88 h). a review of the multi-wavelength properties of ae aqr is provided by meintjes, odendaal and van heerden elsewhere in this volume. it has been shown by meintjes (2002) and schenker et al. (2002) that the properties of this source indicate a high mass transfer history, during which ae aqr could have been a supersoft source. during this phase, accretion disc torques would have been able to spin-up the wd to such a short rotation period. the probable association of the ∼67 s period in cal 83 with a short wd spin may place this source on a similar evolutionary path as ae aqr. evidence indicating that cal 83 has already passed through quite a long mass transfer history does indeed exist. the high wd mass, just below the chandrasekhar limit, is one of the factors supporting an extended period of mass accretion. however, with the wd being more massive than the wd in ae aqr, it may be driven over the chandrasekhar limit before entering the cv stage. emission lines of ionized carbon, nitrogen and oxygen are present in the optical and uv spectra of cal 83, and as remarked by cr87, the low ratio of h to he ii and cno lines indicate a donor with an h-poor envelope. therefore the secondary may already have shed most of its envelope during a long period of mass transfer to the wd, with cno cycling resulting in the observed line ratios. assuming that the secondary mass is as low as ∼0.61 m� (see §2.2), the roche lobe of the secondary should be r2 ∼ 1.7 r�. as the secondary has to be in contact with its roche lobe for roche lobe overflow, this implies a donor that is too large for its mass, possibly an evolved star that has lost a large fraction of its envelope, also supporting the long mass transfer history. the thermal time-scale is of the order τth ≤ 5 × 107 (m2/0.61 m�) −2 (r2/1.7 r�) −1 yr (e.g. meintjes 2002 and references therein), while the spin-up timescale corresponding to a ∼1.3 m� wd accreting from a disc extending to its surface is τs−u ≤ 107 yr (e.g. wang 1987). these time-scales are comparable, showing that the wd in cal 83 could have been spun-up to a short spin period if it has been in the supersoft phase with its characteristic high mass transfer rates for a sufficiently long time. to sustain a τ ∼ 107 yr mass transfer, q > 5/6 is required for a conservative process, as described by the vdh92 model. however, evidence indicating significant mass and accompanied angular momentum losses is present in the optical spectra, suggesting a nonconservative mass transfer process. given the possible wd spin period of ∼67 s, the outflow mechanism 96 supersoft x-ray source cal 83: a possible ae aqr-like system may be the ejection of accreting material by a similar magnetospheric propeller mechanism as in ae aqr. for non-conservative mass transfer, the roche lobe will shrink according to the relation ṙl,2 rl,2 = − ṁ2 m2 [ 5 3 − 2 (1 −α) q − 2 3 qα 1 + q − 2η m2 (gmwdrcirc) 1/2 jorb ] + 2j̇ j , (1) where α and η represent the fraction of mass and angular momentum transferred through l1 that is lost from the system (e.g. meintjes 2002). therefore, for cal 83, mass transfer that can sustain nuclear burning is perhaps only possible in the phase when the evolved donor is expanding, with significant angular momentum loss contributing substantially in keeping the roche lobe in contact with the donor. this will be explored more quantitatively in the future. 4 conclusion it has been shown that cal 83 exhibits a transient ∼67 s x-ray periodicity. this is expected to be associated with the wd spin period, which would make it one of the shortest known spin periods in the white dwarf binary population. results from optical data support the scenario of accretion through an accretion disc, creating the possibility of the wd being spun-up by disc torques during an extended period of mass accretion, similar to what happened during the evolution of ae aqr. optical spectra also present evidence of mass outflows in cal 83, which can be related to the ejection of accreting blobs from a fast rotating wd magnetosphere or accretion disc winds. however, many questions about this fascinating source still remain unanswered, and detailed follow-up studies are essential. acknowledgement the authors would like to thank the conference organisers for the invitation to present this work. the observations reported here were obtained with xmmnewton, salt (proposal 2012-2-rsa uksc-006), and the saao 1.9-m telescope. the authors would also like to thank the salt team for performing the observations, and marissa kotze for technical support and the use of the shoc pipeline. iraf was also used during the optical data analysis. the financial assistance of the south african ska project towards this research is hereby acknowledged. references [1] cowley a. p., schmidtke p. c., crampton d., hutchings j. b., 1998, apj, 504, 854 [2] crampton d., cowley a. p., hutchings j. b., schmidtke p. c., thompson i. b., liebert j., 1987, apj, 321, 745 [3] eracleous m., horne k., 1996, apj, 471, 427 [4] fender r. p., southwell k., tzioumis a. k., 1998, mnras, 298, 692 doi:10.1134/1.1562213 [5] ibragimov a. a., suleimanov v. f., vikhlinin a., sakhibullin n. a., 2003, astronomy reports, 47, 186 [6] kahabka p., van den heuvel e. p. j., 2006, super soft sources. cambridge university press, new york, pp 461–474 doi:10.1086/192065 [7] kaitchuck r. h., schlegel e. m., honeycutt r. k., horne k., marsh t. r., white ii j. c., mansperger c. s., 1994, apjs, 93, 519 doi:10.1086/426382 [8] lanz t., telis g. a., audard m., paerels f., rasmussen a. p., hubeny i., 2005, apj, 619, 517 doi:10.1046/j.1365-8711.2002.05731.x [9] meintjes p. j., 2002, mnras, 336, 265 doi:10.1093/mnras/stt2111 [10] odendaal a., meintjes p. j., charles p. a., rajoelimanana a. f., 2014, mnras, 437, 2948 doi:10.1093/mnras/stt645 [11] rajoelimanana a. f., charles p. a., meintjes p. j., odendaal a., udalski a., 2013, mnras, 432, 2886 doi:10.1046/j.1365-8711.2002.05999.x [12] schenker k., king a. r., kolb u., wynn g. a., zhang z., 2002, mnras, 337, 1105 [13] schmidtke p. c., cowley a. p., 2006, aj, 131, 600 [14] van den heuvel e. p. j., bhattacharya d., nomoto k., rappaport s. a., 1992, a&a, 262, 97 [15] wang y.-m., 1987, a&a, 183, 257 discussion mariko kato: how do you determine the secondary mass for cal 83? is it really so small? i expect q ≥ 1 to explain anti-correlation with x-ray. alida odendaal: the radial velocity semiamplitude of the he ii λ4686 line was used to calculate 97 http://dx.doi.org/10.1134/1.1562213 http://dx.doi.org/10.1086/192065 http://dx.doi.org/10.1086/426382 http://dx.doi.org/10.1046/j.1365-8711.2002.05731.x http://dx.doi.org/10.1093/mnras/stt2111 http://dx.doi.org/10.1093/mnras/stt645 http://dx.doi.org/10.1046/j.1365-8711.2002.05999.x a. odendaal et al. m2 ∼ 0.61 m�, which is only valid if the he ii emission is uniformly distributed around the wd. this assumption can be highly problematic (e.g. kaitchuck et al. 1994), therefore this approach does not yield a conclusive m2. a more reliable wd radial velocity profile may be obtained by constructing doppler tomograms for the emission lines. margaretha pretorius: is there a radio detection of cal 83? alida odendaal: no, but an upper limit of < 0.12 mjy at 3.5 and 6.3 cm was determined by fender, southwell and tzioumis (1998) from data obtained with the australia telescope compact array. 98 introduction cal83: observations and results the 67 s x-ray pulsation optical spectroscopy with salt fast photometry with shoc the evolution of cal83: an ae aqr analogue? conclusion 46 acta polytechnica ctu proceedings 2(1): 46–49, 2015 46 doi: 10.14311/app.2015.02.0046 transient processes in a binary system with a white dwarf. d. a. kononov1, d. v. bisikalo1, v. b. puzin1, a. g. zhilkin1 1institute of astronomy of the ras, moscow, russia corresponding author: dkononov@inasan.ru abstract using the results of 3d gas dynamic numerical simulations we propose a mechanism that can explain the quiescent multihumped shape of light curves of wz sge short-period cataclysmic variable stars. analysis of the obtained solutions shows that in the modeled system an accretion disk forms. in the outer regions of the disk four shock waves occur: two arms of the spiral tidal shock; “hot line”, a shock wave caused by the interaction of the circum-disk halo and the stream from the inner lagrangian point; and the bow-shock forming due to the supersonic motion of the accretor and disk in the gas of the circum-binary envelope. in addition, in our solutions we observe a spiral precessional density wave in the disk. this wave propagates from inside the disk down to its outer regions and almost rests in the laboratory frame in one orbital period. as a results every next orbital period each shock wave passes through the outer part of the density wave. supplying these shocks with extra-density the precessional density wave amplifies them, which leads to enhanced energy release at each shock and may be observed as a brightening (or hump) in the light curve. since the velocity of the retrograde precession is a little lower that the orbital velocity of the system, the same shock wave at every next orbital cycle interacts with the density wave later than at the previous cycle. this causes the observed shift of the humps over binary phases. the number of the shock waves, interacting with the density wave determines the largest number of humps that may be observed in one orbital period of a wz sge type star. keywords: cataclysmic variables dwarf novae optical spectroscopy photometry numerical simulations individual: ss cyg ≡ bd+42◦ 4189a. 1 introduction wz sge stars (subclass of su uma stars) are deeply evolved close binary systems with very short orbital periods (∼ 80) min and low mass ratios of the components (q < 0.1). as a rule these objects are very faint (fainter than 15m). the most remarkable observational feature of these stars in quiescence are their multi-humped orbital light curves. usually astronomers call them double-humped.there are a number of models proposed to explain a double-humped shape of the light curves of wz sge stars. one of them, advocated by aviles et al. (2010), explains the formation of the humps by visibility conditions of two arms of the tidal spiral shock caused by the action of the 2:1 tidal resonance. another model, mentioned, for instance, by wolf (1998) and silvestri et al., (2012) supposes that the stream from the inner lagrangian point may overflow the disk after it ricochets on the disk rim and forms an additional hot spot at the opposite side of the disk, i.e. two hot spots give two humps in the light curve. these two models would have worked perfectly had not there been a number of observational facts that cannot be explained by them. we, above, deliberately called the light curves of wz sge stars “multihumped” instead of commonly used “double-humped”, since there are a number of works where observers report from one to four humps in one orbital period (see, e.g., araujo-betancor et al. (2005), katysheva & shugarov (2009), kitsionas et al. (2005), and pavlenko (2009)). besides, as shown in the mentioned papers, even in the same system the number of the humps varies depending on the epoch of observations. thus, only two bright sources, implied by the models, described above, are not enough to explain more than two humps. one more feature of the humps that cannot be explained by, say, the visibility conditions of the arms of the tidal shock is that they tend to shift over binary phase when observed in time-distant nights. the point is that the positions of the tidal shock arms in an accretion disk are determined by the roche potential and fixed in the coordinate frame, rotating along with the binary. thus, their contributions must always be observed at the same binary phases in every night. it is commonly accepted that the most important observational manifestations of cvs and, in particular, wz sge stars, come from gas dynamical processes initiated by the mass transfer between the components. 46 http://dx.doi.org/10.14311/app.2015.02.0046 transient processes in a binary system with a white dwarf. therefore, in order to understand what happens in wz sge stars and what structures and processes contribute to their observational manifestations one needs to study gas dynamics of these objects. in this paper we, by means of 3d gas dynamic simulations, aim to investigate the flow structure in a representative of wz sge stars and propose a physical model, capable to explain the multi-humped nature of their light curves. in section 2 we describe the simulations and show their results. in section 3, based on the results of gas dynamic simulations we consider a physical model of the formation of the humps in orbital light curves of wz sge stars. in section “conclusions” we summarize our results and discuss further studies. 2 numerical simulations to investigate flow structure in wz sge stars we have performed gas dynamic simulations of a close binary system having the parameters of v455 and. this system is one of the brightest representatives of the class and has well defined parameters (araujo-betancor et al. (2005)) that are as follows. the secondary mass is msec = 0.07m�, and the primary mass (white dwarf) is mwd = 0.6m�. the binary separation of the system is a = 0.56r�. the mass transfer rate in the system is estimated as ṁ = 10−11m�/year. the orbital period of the system is porb = 81.08 min. to model the flow structure in the v455 and system we use a numerical hd/mhd code “nurgush” (zhilkin & bisikalo (2009)). when modeling, we set the magnetic field of the accretor equal to zero, so this is a pure gas dynamic case. the simulations are conducted until the system enters a quasi-stationary regime, i.e. the total mass of the disk varies by less than 2% in one orbital period. approaching this regime takes approximately 30 orbital periods. the results of the calculations are shown in fig. 1. in its panels we show density distributions in the equatorial plane of the system for two time moments τ = 33.53porb and τ = 34.13porb. analysis of the obtained results shows that an elliptic accretion disk forms in the system. one also can see that in the disk a spiral density wave occurs. its approximate location is denoted by the white dashed lines. this precessional density wave, first described by bisikalo et al. (2004), occurs due to the retrograde precession (apsidal motion) of elliptic flow lines, caused by the tidal action of the secondary. the angular velocity of the retrograde precession of this wave (its absolute value) is only a few precent lower than the orbital angular velocity of the system. this means that in the observer’s frame in one orbital period this wave almost rests and, in the frame, rotating with the binary, it makes almost a round during the same period. this effect is wellseen in fig. 1 where at two time moments, separated by approximately 0.6porb, the outer part of the wave occupies almost opposite locations. figure 1: density distributions in the equatorial plane of the system for two time moments τ = 33.53porb (top panel) and τ = 34.13porb (bottom panel). the white dashed line denotes the position of the precessional density wave. 3 model of the multi-humped light curve in fig. 1 one can clearly see that the precessional spiral density wave in our solution propagates down to the outer regions of the accretion disk. from our previous studies (see, e.g., bisikalo & kononov (2010), kononov et al. (2012)) we know that in the outer regions of the disk four most powerfull shock waves should form. these are: two arms of the spiral tidal shock; “hot line”, a shock wave caused by the interaction of the circumdisk halo and the stream from the inner lagrangian point; and the bow-shock forming due to the supersonic motion of the accretor and disk in the gas of the 47 d. a. kononov et al. circum-binary envelope. since, as we noted above, the precessional density wave rests in the observer’s frame and propagates down to the outer regions of the disk, the latter four shocks must pass through its outer part in every next orbital period. when a shock wave passes through the outer part of the density wave it is supplied with extra-density. this must cause amplification of the shock wave. indeed in fig. 2, where we plot pressure distributions for the time moments, corresponding to those of fig. 1, we can see that in the location of the outer part of the density wave pressure increases. the cases shown in fig.2 correspond to time moments when the density wave interacts with the arms of the tidal shock. in the upper panel of fig. 2 the density wave interacts with the closer-to-donor arm of the tidal shock. note that at this moment the pressure in the location of the opposite arm is remarkably lower. in the lower panel the farther-from-donor arm of the tidal shock is more pronounced, while the opposite arm is “inactive”. summarizing the above behavior, we can propose the following model that can explain the formation of the humps in quiescent orbital light curves of wz sge stars. in a system an elliptic accretion disk forms. under the tidal action of the secondary, in this disk, a precessional density wave develops. this wave propagates down to the outer regions of the disk. since the angular velocity of the retrograde precession of the wave is only a few percent lower than the orbital angular velocity of the system, the density wave almost rests in the observers frame. thus the four main shock waves also located in the outer regions of the disk consecutively pass through the outer part of the density wave every next orbital period. when the shock wave interacts with the density wave it is supplied with extra-density and, hence, amplified. the amplification of the shock wave must increase the energy release at the shock that should be seen as a hump in the light curve of the system. we should note that our model can explain not only the origin of the humps but, as well, their behavior. in particular, in the frame of our model, we can easily explain why the humps shift over binary phases from night to night. the point is that the angular velocity of the retrograde precession is not exactly the same as the orbital velocity of the system, it is a little lower. thus, at the next round a shock wave approaches the outer part of the density wave a little later than at the previous one and the interaction between the two waves starts later. the corresponding humps, hence, must shift forward over the binary phase. one more feature of the wz sge light curves that should be explained is the varying number of the humps. we show above that in principle one can observe not two but up to four humps in the light curve, since the density wave may interact with four most powerful shock waves located in the outer regions of the disk. however, the number of seen humps may depend on the visibility conditions of the amplified shock. for example, if a shock is amplified when it is situated at the moment of observations at the opposite from the observer side of the disk, radiation from the shock must pass through the disk and, in principle, may be reduced. besides, the shocks may be of different power. in order to have a clear answer to this question we need to have more observations. figure 2: pressure distributions in the equatorial plane of the system for two time moments τ = 33.53porb (top panel) and τ = 34.13porb (bottom panel). the black dashed line denotes the position of the precessional density wave. 4 conclusions we have performed 3d gas dynamic simulations of the flow structure in a short-period cataclysmic variable star of wz sge type. analysis of the results has shown 48 transient processes in a binary system with a white dwarf. that in the system an elliptic accretion disk forms. in the disk a precessional density wave develops. the wave propagates down to the outer regions of the disk. bisikalo et al. (2004) showed that the angular velocity of the retrograde precession of this wave is only a few percent lower than the orbital velocity of the system. thus, the precessional density wave almost rests in the observer’s frame. from our previous studies we know that in the outer regions of the disk four shock waves form. every next orbital cycle these shocks consecutively pass through the outer part of the density wave. as a result every shock is supplied with extra-density and amplified. the amplification of a shock causes enhanced energy release and may be seen in the light curve as a hump. since the density wave retrogradely precess with the velocity that is a little lower than the orbital velocity, at every next round the same shock passes through the outer part of the density wave a little later. thus the corresponding hump in the light curve at every next round occurs at a later binary phase. this, in the frame of our model, allows us to explain the observational fact that the humps shift over binary phases when observed at different nights. the variable number of the humps may be explained by the visibility conditions of the amplified shock wave. our studies of wz sge stars continue and with further accumulation of observational results we expect to have a more detailed model of the formation of light curves in these cataclysmic variable stars. acknowledgement this work is supported by the russian foundation for basic research (projects 11-02-00076, 12-02-00047, 1302-00077, 13-02-00939, 12-02-31031), the federal targeted program ”science and science education for innovation in russia 2009–2013”, grant of the president of russia for young scientists (mk-2432.2013.2) references [1] aviles, a.; zharikov, s.; tovmassian, g. at al.: 2010, the astrophysical journal, 711(1), 389-398. doi:10.1088/0004-637x/711/1/389 [2] wolf s., et al.: 1998, a&a, 332, 984. [3] n. m. silvestri, p. szkody, a. s. mukadam et al.: 2012, the astronomical journal, doi:10.1088/0004-6256/144/3/84 [4] araujo-betancor, s.; gänsicke, b. t.; hagen, h.-j. at al.: 2005, a&a, 430, 629-642. [5] katysheva, n., shugarov, s.: 2009, journal of physics: conference series, 172(1), 12-44. [6] s. kitsionas, o. giannakis, e. harlaftis, h., et al.: 2005, the astrophysics of cataclysmic variables and related objects, asp conference series, 330, [7] pavlenko e.,: 2009, journal of physics: conference series, 172(1), id. 012071. [8] a.g. zhilkin, d.v. bisikalo: 2009, asp conference series, 406, 118 123. [9] d. v. bisikalo, a. a. boyarchuk, p. v. kaigorodov, o . a. kuznetsov, and t. matsuda: 2004, astronomy reports, 48(6), 449456. doi:10.1134/1.1767212 [10] bisikalo, d. v., kononov, d. a.: 2010 memorie della societ? astronomica italiana, 81, 187 [11] kononov, d. a., giovannelli, f., bruni, i. and bisikalo, d. v.: 2012, astronomy & astrophysics, 538, id.a94, 7 49 http://dx.doi.org/10.1088/0004-637x/711/1/389 http://dx.doi.org/10.1088/0004-6256/144/3/84 http://dx.doi.org/10.1134/1.1767212 introduction numerical simulations model of the multi-humped light curve conclusions 252 acta polytechnica ctu proceedings 2(1): 252–256, 2015 252 doi: 10.14311/app.2015.02.0252 the recurrent eclipsing nova u sco: a short review r. gonzález-riestra1 1xmm science operations centre, esac, madrid, spain corresponding author: rosario.gonzalez@sciops.esa.int abstract u scorpii is the recurrent nova with the shortest inter-outburst period, only ten years. the last active phase took place at the beginning of 2010, and it provided a large amount of data from both ground-based and space observatories. this paper reviews some of the more relevant recent findings and points out some, still unanswered, questions. keywords: recurrent novae uv x-rays individual: u sco. 1 introduction u scorpii is a recurrent nova that undergoes frequent outbursts. it is also an eclipsing binary (i=83 deg, thoroughgood et al. 2001), hence the geometrical parameters of the system are known with great accuracy. the orbital period is 1.23 days (schaefer 1990). ten outbursts of this system have been observed, the first one in 1863 (schaefer 2010). outbursts take place quite regularly, every ≈ 10 years. there are two long intervals of about 20 years during which no outbursts were observed (around 1927 and 1957), but very likely they were missed because the system was too close to the sun when they took place. the system is characterised by extremely short timescales: the recurrence time is only ten years, the shortest known in this type of objects, the system goes into outburst in a few hours, and decays in a few days, with t2=1.2 days and t3=2.6 days. in accordance with all this, very high velocities of the order of 104 km s−1 have been observed. all the above implies that the primary of the system is a high-mass white dwarf. the nature of the secondary star is still unclear, but it is very likely slightly evolved, k2 iv (anupama and dewangan 2000). there is a general agreement on the sizes and masses of both stars: r2 ≈2.4 r�, m2 ≈1.3 m�, m1 ≥1.37 m� (hachisu et al. 2000, thoroughgood et al. 2001). the size of the orbit is ≈ 6.7 r�, and the distance to the system has been estimated to be 12 kpc (schaefer 2010). 1.1 before 2010: uv data iue observed u sco during the first weeks of the 1979 outburst. the uv spectrum presented at the beginning low ionisation lines, and strong p-cygni profiles in the resonance lines. later on, there was an overall increase in the ionisation level, so that in the last observations figure 1: evolution of the ultraviolet spectra of u sco during active phases. spectra of days 4-16 were taken with iue during the 1979 outburst, while the last two were taken with hst-stis in 1999. 252 http://dx.doi.org/10.14311/app.2015.02.0252 the recurrent eclipsing nova u sco: a short review the strongest lines corresponded to he ii, n v, c iv, n iv, and possibly [mg v] (williams et al. 1981, barlow et al. 1981). it is worth noting that the he ii 1640 å line was always much narrower than the others. further uv observations were obtained with hststis during the first two months of the 1999 outburst, confirming the evolution observed before (fig. 1). the last spectra showed only high ionisation broad lines and a narrow he ii line. 1.2 before 2010: x-ray data u sco was known to be a strong x-ray emitter already before the 2010 outburst. the sax observation on day 19 of the 1999 outburst, when the system was detected as a supersoft source, is described by kahabka et al. (1999). these authors fitted a model atmosphere to the sax spectrum, finding a temperature of 9×105 k and a luminosity in the range 2-20×1037 erg s−1. they also found evidence for a weak harder component, and reported variations of the order of 50% during the observation. 2 the 2010 outburst by the beginning of the 21st century the recurrence time of u sco was well established, 10±2 years (schaefer 2005). the last outburst had taken place in 1999, and the date of the next one was predicted to be 2009.3±1. in very good agreement with this prediction, the outburst took place on january 28 2010 (2010.08). the time to rise to maximum brightness was only 9 hours. the optical light curve was very complex, with many different features: there was an early decline phase, a plateau, a second decline and a second plateau. and superimposed on all this, there were short-lived phenomena: flares, dips and the resumption of eclipses (see a detailed description in schaefer et al. 2011). 2.1 x-ray observations 2.1.1 swift the 2010 outburst of u sco was followed by all the x-ray missions, swift, suzaku, chandra and xmmnewton. particularly relevant is the monitoring carried out by swift since, apart from its own scientific value, it allowed to schedule efficiently target of opportunity observations with the other missions, as described below. swift observed u sco daily during two months. supersoft x-ray emission was detected around day 12, when the flux started to increase until reaching a maximum around day 31, to decrease slowly afterwards. the appearance of the supersoft emission roughly coincided with the first extended optical plateau, while the x-ray turn-off started at the time of the second optical decline (schaefer et al. 2010). 2.1.2 suzaku suzaku took three observations of u sco, but only the last one, on day 15, provided useful data, and a surprising result: the detection of a shallow x-ray eclipse (with a depth of ≈30%) in coincidence with the predicted time of the optical eclipse. from the duration of this eclipse, takei et al. (2013) estimated the size of the x-ray emitting region in 5 r�. 2.1.3 high resolution spectroscopy: chandra and xmm-newton a further x-ray observation was taken on day 18, this time with chandra letg. unfortunately, it was short, and taken out of eclipse, and therefore the suzaku result could not been confirmed. orio et al. (2013) claimed that the deep absorption lines seen in the letg spectrum (see top panel of fig. 2) are not p-cygni profiles, but just the superposition of blue-shifted absorptions from the white dwarf atmosphere (whose shift is due to a wind) and the emission lines from the ejecta. these authors derived a temperature of 7×105 k and a luminosity of 7×1036 erg s−1. this is to be compared with the sax estimations, with a much higher luminosity. it is not clear whether the 1999 outburst was intrinsically more luminous, or if the discrepancy is just due to the different atmosphere models used for both determinations. two more x-ray high resolution spectra were obtained with xmm-newton on days 23 and 35, the second one just after the x-ray peak. the xmm-rgs spectrum of the first observation showed similar emission lines as the chandra spectrum, but no absorptions. this was interpreted by orio et al. (2013) as an indication of the end of the wind phase of the white dwarf. the model atmosphere fits presented by these authors showed an increase in the temperature of the white dwarf from day 18 to day 35, at a a nearly constant x-ray luminosity. the emission lines in the day 23 spectrum can be explained by collisional ionisation, while the situation is more complex on day 35, requiring also the existence of a strong uv field. there are some unidentified lines, and others whose presence just cannot be explained. the three grating spectra are shown in left panel of fig. 2. the x-ray light curves also provided interesting information. the first xmm observation started at orbital phase 0.8, thus a clear x-ray eclipse was expected, but an irregular light curve was seen instead, with dips and flares. these dips are an indication of clumps of 253 r. gonzález-riestra material in the re-forming accretion disk (ness et al. 2012). on the contrary, on day 35 the eclipse is very well defined. it is worth to note that on this date the depth of the eclipse is 50%, showing that the x-ray source is very extended, likely a thomsom scattering corona (ness et al. 2012, orio et al. 2013). takei et al. (2013) interpreted the difference in the x-ray eclipses of days 15 and 35 as a shrinking of the x-ray emitting region, from 5 to 4 r�. the optical eclipse mapping done by schaefer et al. (2011) also indicates a shrinking of the optical source, from 4 r� on day 20 to 2.2 r� on day 50. after day 20, the source of the optical light changes, from spherically symmetric to disk-like, first with a bright rim, later brighter at the centre. xmm-newton obtained uv photometry simultaneously with the x-ray observations. on day 23 there was a very well defined uv eclipse, and the light curve was rather smooth, with some structures, as a flare that can be coincident with a similar feature in the x-ray light curve. the depth of the minima in both uv and x-ray light curves was similar (≈50%). the eclipse on day 35 is twice as deep in the uv than in x-rays, indicating a smaller emitting area (ness et al. 2012). figure 2: x-ray high resolution spectra (left panel) and light curves (right panel) of u sco during the 2010 outburst. from top to bottom, chandra letg data taken on day 18, and xmm-newton rgs spectra and epic light curves obtained on days 23 and 35. 2.2 optical spectroscopy also interesting to mention is the evolution of the optical spectra. mason et al. (2012) reported the appearance of narrow emission lines of h, he i and he ii on day 8. the radial velocity curve of these lines (that does not mach either of the two stars) indicate that they form in an accretion flow suggesting that the disk was already re-established. mason (2011) studied the abundance in the ejecta and found a high ne/o ratio, concluding that the white dwarf belongs to the o-ne-mg class, with the implications for the further evolution of the system. finally, in this outburst, forbidden emission lines have been observed for the first time, around day 46 (mason et al. 2012, dı́az et al. 2010). this result must be taken with caution, as it can be just a selection effect, since this outburst had a much better observational coverage than the previous ones. 2.3 the evolution of the outburst as shown in the previous sections, the last outburst of u scorpii has provided a large amount of data that have given us to a more clear view of the behaviour of the system. 254 the recurrent eclipsing nova u sco: a short review models show he disk was destroyed in the outburst (drake and orlando 2010). but only a few days later there were already evidences of accretion being re-established, e.g.: • appearance of narrow optical lines on day 8, • optical and x-ray eclipses on day 15, indicating the existence of a large emitting region ≈ 4-5 r�, • optical flickering and dips in the x-ray light curve on day 23. analysis of the optical light curve shows that by day 40 the disk was already fully re-established. table 1: summary of the evolution of the 2010 outburst of u scorpii day event 0 disk destroyed 8 he, hei and heii narrow emission lines 9-15 optical flares 12 supersoft phase 15 x-ray and optical eclipses 20 end of the white dwarf mass-loss 23 x-ray dips 24 flickering 30 rim-bright disk 40 optical dips 46 nebular spectrum 3 conclusions this system represents an unique opportunity to study the outburst mechanisms of recurrent novae, as it allows us to compare data obtained in several different cycles, with relatively similar instrumentation. despite of the large amount of available data, or possibly due to that, there are still many open issues. for instance, when going through the literature, the values of the mass of the envelope cover a range of 60, even for the same outburst. williams at al. (1981) and anupama and dewangan (2000) get the same value of 10−7 m� for the 1979 and the 1989 outbursts, respectively. on the other hand, takei et al. (2013) obtain 6×10−6 m� for the last outburst. also striking is the wide range of helium abundance values (n(he)/n(h)), as high as 4.5 (evans et al. 2001), but also as low 0.07 (maxwell et al. 2012). is there is a real difference in the mass ejected in the different outbursts? how can the values of the helium abundance be reconciled? obviously, the methods used to derive this quantities and the underlying assumptions are very different. a deep, uniform analysis of all existing data is necessary to find the reason for these discrepancies and to better understand the behaviour of the system. acknowledgement i warmly thank the organisers for their kind invitation to this meeting. references [1] anupama, g. and dewangan, g.: 2000, aj 119, 1359 [2] barlow, m., brodie, j., brunt, c. et al.: 1981, mnras 195, 61 [3] dı́az, m., williams, r., luna, g. et al.: 2010, aj 140, 1860 [4] drake, j. and orlando, s.: 2010: apj, 720, l195 doi:10.1088/2041-8205/720/2/l195 [5] evans, a., krautter, j., vanzi, l. and starrfield, s.: 2001, a&a 378, 132 [6] hachisu, i.; kato, m., kato, t. and matsumoro, k.: 2000, apj 534, l189 [7] kahabka, p., hartmann, h., parmar, a. and negueruela, i.: 1999. a&a 347, 43 [8] mason, e., 2011, a&a 532, l11 [9] mason, e., ederoclite, a., williams, r. et al.: 2012, a&a 544, 149 [10] maxwell, m., rushton, m., darnley, m. et al.: 2012, mnras 419, 1465 doi:10.1111/j.1365-2966.2011.19803.x [11] ness, j.-u., schaefer, b., dobrotka, a. et al.: 2012. apj 745, 43 doi:10.1088/0004-637x/745/1/43 [12] orio, m., behar, e., gallagher, j. et al.: 2013, mnras 429, 1342 doi:10.1093/mnras/sts421 [13] takei, d., drake, j., tsujimoto, m. et al.: 2013, apj 769, l4 doi:10.1088/2041-8205/769/1/l4 [14] schaefer, b.: 1990, apj 355, l39 [15] schaefer, b.: 2005, apj, 621, l53 doi:10.1086/429145 255 http://dx.doi.org/10.1088/2041-8205/720/2/l195 http://dx.doi.org/10.1111/j.1365-2966.2011.19803.x http://dx.doi.org/10.1088/0004-637x/745/1/43 http://dx.doi.org/10.1093/mnras/sts421 http://dx.doi.org/10.1088/2041-8205/769/1/l4 http://dx.doi.org/10.1086/429145 r. gonzález-riestra [16] schaefer, b.: 2010, apjs 197, 275 doi:10.1088/0067-0049/187/2/275 [17] schaefer, b., pagnotta, a., osborne, j. et al.: 2010, atel 2477 [18] schaefer, b., pagnotta, a., lacluyze, a. et al.: 2011, apj 742, 113 doi:10.1088/0004-637x/742/2/113 [19] thoroughgood, t., dhillon, v., littlefair, s. et al.: 2001, mnras 327, 1323 [20] williams, r., sparks, w. gallagher, j. et al.: 1981, apj 251, 221 256 http://dx.doi.org/10.1088/0067-0049/187/2/275 http://dx.doi.org/10.1088/0004-637x/742/2/113 introduction before 2010: uv data before 2010: x-ray data the 2010 outburst x-ray observations swift suzaku high resolution spectroscopy: chandra and xmm-newton optical spectroscopy the evolution of the outburst conclusions 269 acta polytechnica ctu proceedings 1(1): 269–273, 2014 269 doi: 10.14311/app.2014.01.0269 active galactic nuclei: jets as the source of hadrons and neutrinos athina meli1,2,3, paolo ciarcelluti3 1department of physics and astronomy, university of gent, belgium 2ifpa, department of astrophysics, geophysics and oceanography, university of liege, belgium 3web institute of physics, www.wiph.org corresponding author: athina.meli@ugent.be abstract active galactic nuclei are extragalactic sources, and their relativistic hot-plasma jets are believed to be the main candidates of the cosmic-ray origin, above the so-called knee region of the cosmic-ray spectrum. relativistic shocks, either single or multiple, have been observed or been theorized to be forming within relativistic jet channels in almost all active galactic nuclei sources. the acceleration of non-thermal particles (e.g. electrons, protons) via the shock fermi acceleration mechanism, is believed to be mainly responsible for the power-law energy distribution of the observed cosmic-rays, which in very high energies can consequently radiate high energy gamma-rays and neutrinos, through related radiation channels. here, we will focus on the primary particle (hadronic) shock acceleration mechanism, and we will present a comparative simulation study of the properties of single and multiple relativistic shocks, which occur in agn jets. we will show that the role of relativistic (quasi-parallel either quasi-perpendicular) shocks, is quite important since it can dramatically alter the primary cr spectral indices and acceleration efficiencies. these properties being carried onto gamma-ray and neutrino radiation characteristics, makes the combination of them a quite appealing theme for relativistic plasma and shock acceleration physics, as well as observational cosmic-ray, gamma-ray and neutrino astronomy. keywords: active galactic nuclei jets relativistic shocks cosmic rays acceleration gamma-rays neutrinos. 1 introduction cosmic rays (crs) are subatomic particles (e.g. protons) and radiation of extra-terrestrial origin. their energies range between a few ev to hundreds of pev. these particles register a power-law spectrum, and the low-energy ones are believed to originate from galactic, while the crs with higher energies originate from extra-galactic sources. specifically, crs above energies of ∼ 1017ev, almost certainly originate from extragalactic sources such as active galactic nuclei (agn) and their relativistic jets. it is the particle shock acceleration in the superalfvenic jet plasmas, which is believed to be the main mechanism responsible for the production of the nonthermal crs (e.g. krymskii, 1977). especially, the importance of the very high energy cr acceleration is a favourable theme of study, as it is also the tev photons and neutrinos theme, which are produced either by the primary proton-photon or primary proton-nucleon interactions, offering further insights about the environment of a considered astrophysical source. agn sources are actually a compact region at the centre of a galaxy, which is believed to be a supermassive black hole. the radiation from an agn is believed to be mainly a result by the accretion of mass around the black hole. the accretion of matter around the spinning core of an agn produces twin, highly collimated fast outflows of hot gas, called jets, extending to distances of a few parsecs to thousands of parsecs. it is of importance to mention that the physics of jet mechanism and the jet composition on very small scales, are not perfectly understood at present. we know thought, that the agn black holes and their jets radiate in all wavebands, from the radio up to the gamma-ray regime via mostly the synchrotron and the inverse-compton scattering processes. moreover with a typical agn luminosity of l ∼ 1042 − 1047 erg/s, it is possible to explain the origin of the very high energy cr flux, under the assumption that these objects follow the star-formation-rate. for example, moran et al. (2001) find an average luminosity from an estimated 95 x-ray active agn within 60 mpc of 4.8 × 1044 erg/s, this distance being the cr absorption horizon at about 2×1020 ev. within the regions of local space accessible to cr diffusion, the energy supply over a hubble time is 6.6 × 10−16 erg/cm3. by comparison, the gamma ray bursts supply is 6.7×10−20 erg/cm3. thus, agn as the most luminous permanent sources in the sky can stand as the most likely candidate for the acceleration of very 269 http://dx.doi.org/10.14311/app.2014.01.0269 athina meli, paolo ciarcelluti high energy crs. interpretation of data on the electron synchrotron radiation observed in the radio regime, suggests the presence of mildly-relativistic shocks with mean boost factors of γ = 10 − 30 in the jets of agn, see biermann and strittmatter (1987). moreover, assuming that hadrons (as we do in this work) are also accelerated along with the electrons in the jet, justify agn to be as good candidates as to provide at least a significant fraction of the extragalactic component of the hadronic cr flux observed. here we will discuss the primary particle shock acceleration mechanism in agn jets, in the context of comparative test-particle simulation studies, for relativistic single and multiple quasi-parallel and quasiperpendicular shocks. we will also briefly highlight the possible consequent tev-photon and neutrino emission. 2 single-relativistic shocks here we will review shock acceleration simulations, based on the relativistic single-shock, test-particle monte carlo code of meli and quenby (2003a,b), as a comparative tool to the multiple-shock acceleration study of the following section. in our simulations we inject a large number of particles n = 104 upstream towards a single planar shock, with an initial particle energy γ = γsh + 5. we use shock lorentz factors of γ = 10 and 30, and we numerically let the particles’ guiding-centers to scatter with the assumed background magnetic field irregularities, around a shock region of a length d = 2 pc. for details on the code see meli and quenby (2003a,b) and meli et al. (2008). briefly, our exemplary simulations show that : the relativistic quasi-parallel (e.g. ψ = 25o) and quasiperpendicular (e.g. ψ = 75o) single-shock acceleration mechanism manifests into a clear deviation of the cr spectral index (compared to standard, unaffected non-relativistic spectral index of ∼ 2.0) with an absent ’universal’ power-law spectrum, which depends on three parameters: i) speed of the shock, ii) shock inclination, and iii) different scattering modes, see table 1. we clearly see a high acceleration efficiency (in terms of emax) for the cases of quasi-parallel shocks. albeit, quasi-perpendicular shocks result into more ’conservative’ acceleration efficiencies and lower attained energies. it is important to note that it is been reported that the lower the turbulence, in terms of the mean-free-path η = λ/rg, the steeper the particle spectrum (meli et al. 2008). near-perpendicular shocks strongly appear to give the steepest spectra compared to quasi-parallel shock cases, and are the less efficient accelerators, see also e.g. niemec and ostrowski (2004), stecker et al. (2007), etc. summarizing, it is widely accepted that (with a few exceptions) as soon as an astrophysical shock speed becomes comparable to the speed of light, the power-law index of the accelerated cr spectra, depends strongly not only on the speed of the shock but also on the shock obliquity (scattering mode is also important, see meli, 2011), identifying clearly flatter spectra for small shock inclinations, with greater acceleration efficiencies. table 1: examples of simulations for single shocks with different shock lorentz factors (γ) and shock inclination angles (ψ), and the attained spectral indices (σ), and maximal particle energies e, respectively. γ ψ = 25o ψ = 75o 10 σ = 2.0,e = 109.3gev σ = 2.4,e = 105.6gev 30 σ = 1.9,e = 109.8gev σ = 2.3,e = 106.0gev 3 multiple-relativistic shocks after the brief insights on the relativistic single-shock acceleration above, we will now discuss a case study of particle acceleration in multiple-shocks, based on the fully relativistic multiple-shock, test-particle monte carlo code by meli and biermann (2013). in the past it was shown that for an infinite number of subsequent non-relativistic shocks with injection at each shock, the flattening of the spectrum (compared to a single shock) extends even to the highest energy particles with a momentum dependence of f(p) ∝ p−3 (e.g. white 1985). here we will investigate the acceleration mechanism in relativistic, multiple, quasi-parallel and quasiperpendicular shocks. we can envisage blobs of hot plasma ejected by the black hole of an agn that can travel along the jet and pass through a sequence of shocks, as it has been also observed in blac pks 1510089 (marscher et al. 2010) and other sources. here we numerically calculate the scattering of the particles’ guiding-center with pre-determined background magnetic field irregularities, around multipleshock regions and within a distance d = 5 pc. at the beginning of the acceleration we allow a single initial injection of particles of a fixed number n = 104, with an initial γ = γsh + 5, and shock lorentz factor γ = 30 (for the first two shocks) and γ = 10 (for the second pair), justified by the decompression-compression conditions of the downstream plasma (melrose and pope 1993) of the precedent shock pattern (i.e., a velocity compression ratio r=3). that is, during the acceleration we allow a probability of escape pesc = 0.1 between shocks due to decompression effects. the escape probability gives the fraction of particles of the downstream distribution of each shock, that will not be accelerated 270 active galactic nuclei: jets as the source of hadrons and neutrinos more. these particles remain in the system and contribute directly to the downstream distribution of the 4-th shock. this factor would naturally result in steeper spectral slopes comparing to a single shock acceleration spectrum as the relativistic conditions render the spectra flatter. assuming also a field compression ratio rbc equal to , and an assumed magnetic field strength of the order of a few mgauss, the distance between each shock is safely set as d = 1pc. we study two cases of shock patterns: i) a set four quasi-parallel shocks (with the same apex to the assumed jet axis), and inclination angles ψ1,ψ2,ψ3,ψ4 = 25o, and ii) a set of four quasi-perpendicular shocks and inclination angles ψ1,ψ2,ψ3,ψ4 = 75 o. table 2: aaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaa γ ψ σ e[gev ] 30 25o 2.6 105.9 30 25o 2.2 106.9 10 25o 2.1 108.6 10 25o 1.5 109.5 table 3: in the tables above we see indicative simulation numbers for two sets of four-shocks, their maximal attained energies e at each shock downstream, and their spectral indices σ. shock inclination angles are denoted as ψ, and γ is the shock lorentz factor. γ ψ σ e[gev] 30 75o 2.7 104.6 30 75o 2.4 105.2 10 75o 2.3 105.9 10 75o 2.2 106.4 briefly, from the tables above, one notices that (i) the spectra by each shock in every shock-set become gradually flatter, following on one hand the findings of single relativistic shocks discussed in the previous section, and on the other hand the trends reported in e.g. melrose and pope (1993) and gieseler and jones (2000) for multiple shocks. (ii) the flatness of the spectra is more evident for the consecutive quasi-parallel shocks than for the quasi-perpendicular ones, following, as expected, the individual shock effects we discussed in the previous paragraph. (iii) the acceleration in the case of quasi-parallel shocks seems very efficient and comparable to single shock acceleration. 4 neutrino and gamma-ray astronomy over the years many models have been created (e.g. stecker, 2005), assuming protons are accelerated at the shocks of agn jets and interact with different photon fields to produce high energy gamma-rays and neutrinos. the cr acceleration cases shown here, have implications for neutrino and gamma-ray astronomy, and especially for the tev sources. we know that only from e.g. the p−γ interaction (where the optical depth must be equal or larger than 1), the consequent decay of the π0 results into a high energy photon (in 0.33 of the cases). also π± particles produce neutrinos. specifically, the probability for π+ production is 0.66. around 0.20 of the proton energy goes into pion, and the products carry equal energy, so that a fraction of 0.25 (0.50) is transferred to each neutrino (photon) and furthermore, one must also consider that neutrinos oscillate. therefore the high-energy photon and neutrino spectra can be connected as dnν deν = 1 8 · dnγ deγ . additionally it can be assumed that the neutrino carries about 1/20th of the proton’s energy (see halzen, 2008). all these can be very helpful in understanding crs and related consequent emissions. nevertheless, in many agn the photon signal at tev energies is not unique, since leptonic processes like inverse compton scattering could also contribute at the same energies. but there is a way of identifying hadronic interactions in astrophysical shocks by estimating the flux of neutrinos. not to let unmentioned, one can also assume that the spectral indices for the tev photon and neutrino emission at a break threshold energy, generally at around 102gev, can be considered to be about the same. it is accepted that not all agn sources can be seen or inferred emitting crs, and this due to either the high galactic background absorptions, or particle’s random diffusion in magnetic fields, or even insufficient power to accelerate very high crs. nevertheless, neutrino signals, could reveal the possible hadronic nature of a tev source. we saw in this work that relativistic quasiperpendicular shocks in agn, single or multiple, cannot accelerate crs to very high energies. we know also that a photon field in the jet of an agn can range from low energies to many tev. we also know (e.g. becker et al. 2011) that the proton energy necessary to produce a delta resonance in the observer’s frame, can be given as, ep ≥ γ 2 (1+z)2 m2∆−m 2 p 4 · e−1γ , where m∆ is the mass of the delta resonance, mp is the proton-mass, eγ is the characteristic photon energy and z is the redshift of the source (e.g. cen a, z = 0.001). therefore, simplistically, one can assume a photon field (i.e. but dense enough for photo-hadronic interactions) to make 271 athina meli, paolo ciarcelluti an estimation of the maximal energy of the proton (as shown in this work) necessary to produce tev-photons and gev-tev neutrinos, or vice-versa, and make further estimations on the correlations of different wavelengths from the same source. as an example here for the case of the favourable tev emitter, cen a, with z = 0.001 and an assumed shock(s) lorentz factor of γ = 30, the proton energy sufficient to produce a delta resonance and with these, tev-photos and neutrinos, can be given as ep ≥ 105 · (1 + z)−2 · ( γ 30 )2 · ( eγ 1 mev )−1 gev . thus, for quasi-perpendicular shocks, the proton spectra found in our simulations reach low energies ep ∼ 106 gev, which could not be directly observed. in order though to achieve tev photons and along with them neutrinos, one would require photon field energies at the jet of equal or greater than eγ ≥ 10 mev. of course these energies can be produced via the inverse compton scattering of synchrotron or external photons with the accelerated electrons certainly present at the jet. therefore one would expect a high energy signal of neutrinos, if the source is of a hadronic nature. what one can in generally conclude from the above brief discussion, is that the extragalactic tev sources, which can not be observed as candidates of ultra-high energy crs, may nevertheless emit high energy neutrinos implying the hadronic nature of the cosmic rays. this fact can be used as a supplementary tool to filterout neutrino source candidates. calculations on flux and further details on radiative consequences are under away. 5 conclusions the relativistic shocks in the jets of agn are a favourable accelerator of high-energy as well as lowenergy crs, via the shock acceleration mechanism. here, we performed comparative monte-carlo simulation studies for single and multiple relativistic shocks, assuming protons as the primary accelerated particle and we showed that relativistic (quasi-parallel either quasi-perpendicular) shocks are very important since they can alter dramatically the cr spectral indices and acceleration efficiency. these properties are carried onto gamma-ray and neutrino emission from the same sources. thus, assuming a hadronic primary accelerated scenario, one could have an additional understanding regarding expected neutrino fluxes from tev sources. specifically, agn with tev gamma-ray emission, and non-observable or low-energy crs, could have the potential to carry an additional high energy neutrino flux signal, mimicking accordingly the primary accelerated particle spectra in the jet, separating it from the inverse compton leptonic or other related processes. accurate flux expectancies of gamma-rays and mostly neutrinos within earth neutrino-observatory potentials (e.g. icecube, km3net), is the next step for deeper understanding of agn jets. acknowledgement a.m. would like to thank the organisers for the invitation to give an advanced talk on the topic of shock acceleration and cosmic rays. she also likes to thank p. l. biermann, j. j. quenby and j. k. becker for contributing their valuable expertise to the cited studies above; she acknowledges fruitful discussions with antares and icecube/icetop collaborations; and she is constantly indebted to a. mastichiadis, s. dimitrakoudis, r. protheroe, a. donea, t. kneiske, m. dieckmann, a. reimer, m. ostrowski, i. brancus, j. wentz, m. giller, a. wolfendale, t. stanev, k.nishikawa for offering, in different occasions over the last years, valuable advice and insights on related studies. references [1] becker, j., meli, a. & biermann p. l., 2011 nucl. ist. meth., 630 269 doi:10.1016/j.nima.2010.06.082 [2] biermann, p. l. & strittmatter, p. a., 1987 apj, 322, 643 [3] gieseler, j. & jones, t. w., 2000, aa 357, 1133 [4] halzen, f., j., 2008 phys. conf. ser., 120 (6) 062004 doi:10.1088/1742-6596/120/6/062004 [5] krymskii, g. f., 1977 akademiia nauk sssr, 234, 1306 [6] marscher, a. p., jorstad, s. g., larionov, v. m., et al., 2010 apj, 710l, 126 doi:10.1088/2041-8205/710/2/l126 [7] meli, a. & quenby, j. j., 2003b aph 19, 649 [8] meli, a., becker, j. k., j., quenby, j. j., 2008 a&a 492, 323 [9] meli, a., 2011, astrspsc 7, 287 [10] meli, a & p. . biermann, 2013 a&a in press [11] melrose, d. b. & pope, m. h., 1993 pasau, 10, 222 [12] moran, e. c., kay, l. e., davis. m., filippenko, a. v., barth., 2001 ast.j., 566, l75 [13] niemiec, j. & ostrowski, m., 2004, apj 610, 851 doi:10.1086/421730 [14] stecker f. w., 2005 phrvd, 72j 7301 [15] stecker f. w., baring j., summerlin e. j., 2007 apj, 667l, 29 doi:10.1086/522005 [16] white r. l., 1985 apj 289, 698 discussion pieter meintjes: you mentioned that strong relativistic shocks will be accompanied by turbulence. can you comment on the role that mhd can possibly play in the 272 http://dx.doi.org/10.1016/j.nima.2010.06.082 http://dx.doi.org/10.1088/1742-6596/120/6/062004 http://dx.doi.org/10.1088/2041-8205/710/2/l126 http://dx.doi.org/10.1086/421730 http://dx.doi.org/10.1086/522005 active galactic nuclei: jets as the source of hadrons and neutrinos acceleration process? is it possible that strong mhd turbulence can distort a power-law spectrum into a kind of relativistic maxwellian as a result of acceleration through mhd waves? if this happens, under what conditions? athina meli: the numerical approach of the model presented here is that of the test-particle approach, investigating a collision-less plasma (kinetic theory) and not a collisional plasma (fluid theory). i note that shocks do exist in collisionless media, i.e. without collisional dissipation (resistivity/viscosity). let me also add though that an mhd description of a collisionless plasma it is not entirely wrong to be assumed, because all fluid equations (except closure) are exact conservation laws (and closure is not too sensitive to microphysics at large scales), but there are many other exceptions mainly when large scales are not independent of the small ones with critical limits. for the present case, the mhd turbulence effect connection can be attained at the context of assuming the strength of scattering modes directly in the code (i.e. strong or weak scattering, as large angle diffusion or pitch angle scattering), by relating it to the diffusion coefficient (κ) which is dependent on the level of media turbulence, denoted by the ratio λ/rg. please see detailed descriptions of the particle kinematics in meli and quenby (2003b), meli et al.(2008) and meli (2011). as i discussed here, the type of scattering hence the level of mhd turbulence in the media can alter the produced spectra in the relativistic shock acceleration cases. jim beall: does the spectra you get from these shocks depend on the specific shock structure? athina meli: the shock structure can alter the steepness of the shock, and i deal with this matter in another recent work (meli, 2013) (not presented here, planar shocks only). maurizion spurlo: cosmic rays with energies above 1019 ev, following the auger measurements, seem to be mainly heavy nuclei. can the acceleration model from agn jets you presented, accelerate heavy nuclei instead protons maintaining the same features? athina meli: the relativistic shocks presented here produce spectra mainly below 1019ev , favoured for protons. nevertheless, in some ideal cases one can obtain easily spectra that reach energies of 1020−21ev, as i showed in a related work. this acceleration model can treat the acceleration of heavy nuclei, but the consideration of added losses must be taken into account in the code, which will mostly steepen the spectrum by an order of magnitude or more, as some of my preliminary test-runs have shown. i note here though an important point: the spectra even above energies of 1019ev seem to be a (unknown) mix of protons and heavy nuclei, since pure heavy nuclei population seems highly improbable. 273 introduction single-relativistic shocks multiple-relativistic shocks neutrino and gamma-ray astronomy conclusions 297 acta polytechnica ctu proceedings 2(1): 297–302, 2015 297 doi: 10.14311/app.2015.02.0297 study of cataclysmic variables with the satellites loft and gaia r. hudec1,2, v. šimon 1,2 1astronomical institute, academy of sciences of the czech republic, cz-25165 ondřejov, czech republic 2czech technical university in prague, faculty of electrical engineering, prague, czech republic corresponding author: rene.hudec@asu.cas.cz abstract the goal of this paper is to discuss the capabilities of esa satellite missions gaia (already in space) and loft (considered for the esa m4 slot) for investigation of cataclysmic variables (cvs). both gaia and loft can contribute to study of cvs and related objects. spectrophotometry and low dispersion spectroscopy are the most important for cv analyses with the observations of gaia. we present the possible strategies for investigation of cvs in the sampled photometric and spectroscopic data provided by gaia. e.g. statistical properties of the long-term activity of various types of cvs can be determined from them. loft can be a promising satellite to provide a sensitive x-ray monitor which will enable to investigate the little studied long-term activity of various types of cvs (especially the magnetic ones) in the x-ray band. keywords: cataclysmic variables dwarf novae intermediate polars x-rays loft gaia. 1 introduction although two esa satellite missions, namely gaia (already in space) and loft (considered for the esa m4 slot), focus on different science areas, both of them can effectively contribute to the investigation of cataclysmic variables (cvs). 2 the esa gaia mission the esa satellite gaia was successfully launched on december 19, 2013. gaia is an ambitious mission to chart a three-dimensional map of our galaxy, in the process revealing the composition, formation and evolution of the galaxy (e.g. de boer et al. 2000 and eyers et al. 2013). gaia will provide unprecedented positional and radial velocity measurements with the accuracies needed to produce a stereoscopic and kinematic census of about one billion stars in our galaxy and throughout the local group. this amounts to about 1 percent of the galactic stellar population. 2.1 astrophysics with gaia motivation of this paper is to outline the possibilities of performing astrophysics of cvs with gaia data. this is not trivial as the main goal of the gaia mission is to create a catalog. the photometric sampling provided by gaia will not be optimal for many variable astrophysical sources. however, the fine spectrophotometry (in reality ultra low resolution spectroscopy) provided by bp/rp photometers will be unique and important for many astrophysical investigations with gaia, including cataclysmic variables science (hudec & šimon 2007a and 2007b; hudec et al. 2012 and 2013). even the sampled gaia data can provide the following properties of the object: 1. determination of amplitude of the brightness variations 2. statistical distribution of brightness 3. absolute magnitude (from distance determined from parallax (and interstellar extinction)). note that these items 1, 2, and 3 even enable to establish the physically justified sequence of types of cvs. 4. color indices and their time variations in principle, it is also possible to search for the periods and cycles of the brightness variations, but one has to very cautious in doing this procedure using the sampled data from gaia. the sampling of the data provided by gaia is not optimal for the astrophysical work; these observations are not dense enough and not equidistantly distributed. typically, one can expect 50–100 photometric points over 5 years. however, additional data can be provided by the ground-based experiments. we propose the following two approaches for classification and verification of cvs in the sampled gaia data. the first one is to search for outbursts or high/low 297 http://dx.doi.org/10.14311/app.2015.02.0297 r. hudec, v. šimon state transitions, i.e. analysis of parameters of the light curve not strongly dependent on sampling. 2.2 statistical properties of the long-term activity of cvs, perspectives of histograms for gaia we investigated the profiles of the light curves of cvs of various types with real observations (daily means) from the afoev database (http://cdsarc.ustrasbg.fr/afoev/). we approximated the gaia sampling of these light curves by selecting the data separated by ∼ 20 days. this procedure yielded the amount of the sampled data and the length of the mapped time segment which roughly corresponded to the expected time of observing with gaia. we investigated which methods can yield correct identification of a cv of a given type almost independent of the sampling. we identified the impact of the gaia data on the investigation of cvs which are very active objects. figure 1: comparison of the observed (one-day means of afoev data (crosses)) and the approximated sampled data of the dwarf nova ah her (closed circles). heavily sampled data are expected from the gaia satellite. notice that it is difficult to resolve the profile of the light curve in the sampled data. the basic profile of the high state and its fluctuations in e.g. novalike systems and supersoft x-ray sources can be plausibly mapped by the sampled gaia data. only the high state/low state transitions may be missed or covered by only one or two data points by gaia. on the other hand, we expect the large outbursts of a dwarf nova to be covered by only a very few data points while short outbursts may be missed in the gaia data set. moreover, the individual outbursts will be probably captured in different phases due to their short duration. gaia will therefore usually not provide us with information about the profile of the outburst. however, we found that the statistical distribution of magnitudes of a given cv is only slightly distorted if a long time (several years) interval is covered. the statistical distribution of brightness and its parameters like the standard deviation, skewness, excess are a representative description of the properties of the long-term activity of cvs. these histograms are only slightly distorted by the sampling of the gaia observations if a long (several years) time segment is covered by the observations. nevertheless, a distortion may appear if the recurrence time of outbursts is close to the sampling time of gaia. figure 2: statistical distribution of the observed data in the dwarf nova ah her (from fig. 1). figure 3: statistical distribution of the approximated sampled gaia data of the dwarf nova ah her (from fig. 1). notice that the histogram is not dramatically influenced by the sampling and is similar to that of the observed data. we conclude that the sequence of the types of cvs, justified by the physics of accretion onto the compact objects, is reflected in the statistical distribution of the long-term brightness variations, hence it is suitable for analysis of the sampled gaia data. the importance of this sequence can be even enhanced if the apparent magnitudes are transformed to the absolute magnitudes by using the distances of cvs determined from the parallaxes. 2.3 rare flares (different from outbursts in dwarf novae) in cvs we investigated the possibility to detect and identify rare phenomena in cvs with gaia. deep monitoring of a large number of objects can lead to discoveries of more cvs with flares (phenomena different from outbursts) like the event discovered by van amerongen 298 study of cataclysmic variables with the satellites loft and gaia & van paradijs (1989) and to detection of additional such events even in the already known systems. we investigated some cvs in bamberg photographic plates which have a similar sampling and coverage as the expected gaia data. we conclude that the discovery of such events is possible with the monitoring planned for gaia (šimon 2010). 2.4 ultra-low dispersion spectra the ultra-low dispersion spectra will be provided by gaia rp/bp photometers. they can play a significant role in (1) searches for prominent spectroscopic variability, and (2) searches for objects with prominent (maybe variable) spectral features. as an example, the spectra of the outburst of v407 cyg (munari et al. 2011) (or of an analogous object) could be investigated by gaia. the spectral appearance of v407 cyg in outburst was a highly peculiar one. the spectrum of this event is completely different from those ever recorded for this object and other symbiotic mira variables. the white dwarf companion to the mira variable experienced an outburst similar to that of classical novae, and its ejecta were moving in the circumstellar environment already filled by the ionized wind of the mira. we summarize several types of cvs and the related systems with very bright emission lines: • supersoft x-ray sources in the optical high states (steiner & diaz 1998). their spectra display intense and broad emission lines (especially hα) (e.g. v sge (herbig et al. 1965)). such systems also often have balmer jump in emission. • classical novae in the nebular phase of the outburst (late phase, weeks to months after the maximum of the optical luminosity) – faint continuum, very strong emission lines of various elements (e.g. munari et al. 2013). • symbiotic systems: very strong emission lines (mainly hα) (e.g. davidson et al. 1978). • dwarf novae: highly variable strong balmer jump (e.g. walker & chincarini 1968). 2.5 the color indices from rp/bp spectra the color indices can be determined from the spectra obtained by the gaia rp/bp photometers. significant results are expected to be immediately available, there is no need to wait for years as in the photometry case. the color indices give important information on the spectral energy distribution. these indices can play an important role in classification and/or verification of cvs. they are also important for a search for the common properties of the sources of a given kind (e.g. to help identify a source as a cv and/or to study the evolution of the spectral profile with time as the cv undergoes various states of its activity). not only the color indices of the object at a given time, but also the time evolution of these indices are therefore important. they can also play a role in resolving among the individual radiation mechanisms (e.g. cyclotron radiation versus thermal emission). even variations of strong emission lines with respect to the continuum can be resolved by the color indices (e.g. hα changes, changes of balmer jump between emission and absorption). the color indices are also important for forming a representative ensemble of events (e.g. outbursts) in a given cv or in a given type of cvs. this is possible also for faint objects detected by gaia (down to mag ∼ 20). the parallaxes from gaia will enable to determine the distances, and hence the absolute magnitudes of many cvs. it will be therefore possible to determine and study the relation between the color indices and the absolute magnitudes of cvs of various types. 2.6 searches for cycles or periods variograms allow us to search for characteristic timescales or quasiperiods which extend just for several epochs of the cycle. variogram characterizes the spatial continuity or roughness of a data set. variograms are important for a search for the superorbital cycles in the long-term activity of cvs and low-mass x-ray binaries because their long-term activity is not periodic. standard period searches therefore often reveal nothing. we used the light curves of cvs of various types for testing. we applied the real observations from the afoev database for feasibility study of gaia data. we approximated the gaia sampling of these light curves by the data separated by ∼20 days. we investigated how the variograms were modified by the sampling. an example, the supersoft x-ray source v sge, is displayed in fig. 4. the amount of the sampled data and the length of the mapped time segment roughly correspond to the expected time of observing of gaia. since the duration of the high or the low state is longer than the time interval between the sampled data, the profile of the light curve is still recognizable. our tests showed that, in the case of the gaia data, variograms are suitable especially for investigation of the types of cvs which display the light curves, whose profiles are gradual. v sge with its alternating high and low states and the amplitude of the brightness variations similar for the individual epochs of the cycle is a good example (figs. 5 and 6). similar results can be obtained for novalike cvs with the episodes of the high and low states. however, it is necessary to be cau299 r. hudec, v. šimon tious in the case of dwarf novae because the outbursts, although with a large amplitude, are usually considerably shorter than the intervals of quiescence. this implies that some outbursts may be missed in the sampled data. it also emerged that binning of the data for variograms (fig. 6) plays a big role. fine binning yields a more precise determination of the cycle-length, but artifacts and false detections can appear, especially if the data are sampled. figure 4: comparison of the observed (one-day means of afoev data (crosses)) and the approximated sampled data (closed circles) of the supersoft x-ray source v sge. heavily sampled data are expected from the gaia satellite. figure 5: variogram for the observed data of the supersoft x-ray source v sge (from fig. 4). n bins refers to the number of bins used for constructing the variogram. the orbital modulation of cvs can be searched for and investigated in the sampled gaia data only under some favorable circumstances. although the length of the orbital period of cv can be considered to be stable (or at most slightly variable) during the proposed lifetime of gaia, it is reasonable to expect that the orbital modulation will be influenced by the long-term variations (e.g. outbursts, episodes of the high and low states). the amplitude of the long-term variations can often be considerably larger than that of the orbital modulation even in the case of a significant amplitude of this modulation (e.g. > 0.8 mag). only if the long-term level of brightness remains almost stable (or if the limited segments of the long-term light curve are selected), the orbital modulation can emerge. a preferable situation will be to fold the gaia data with the already known orbital period, and to investigate if and how the profile of the modulation varies with the changes between the states of activity. the orbital period can be determined for example by the follow-up observations of the ground-based telescopes. this is one of the cases of co-operation of the orbital and the ground-based observing components of the project. figure 6: variogram for the approximated gaia data of the supersoft x-ray source v sge (from fig. 4). a higher value of n bins can yield a better determination of the possible cycle-length, but also a bigger noise. 3 esa loft loft (the large observatory for x-ray timing) is specifically designed to exploit the diagnostics of very rapid x-ray flux and spectral variability in compact objects, yielding unprecedented information on strongly curved spacetimes and matter under extreme conditions of density and magnetic field strength (feroci et al. 2012; brandt et al. 2013). loft is designed to investigate variability from submillisecond quasi-periodic oscillations (qpo) to years long transient outbursts. loft belongs to esa medium class mission (m3 candidate) and was selected in feb 2011 to complete a 3year phase 0/a study. final down selection on jan 2014 (5 competed for single launch slot) was however negative for loft. there is however a chance to be selected for m4. loft is observatory-type mission: there are 2 onboard instruments and science data center provided by the community. the originally expected launch date was 2022–2024 timeframe (soyuz launcher) with 4+1 years mission lifetime and low earth orbit (550 km, 2 ground stations, 7 gb science data/orbit). the loft large area detector (lad) has an effec300 study of cataclysmic variables with the satellites loft and gaia tive area ∼20 times larger than any largest predecessor, uniquely combined with a ccd-class energy resolution. loft lat is suitable for observing cvs in a wide xray band of 1–40 kev, with very fine time resolution. the loft wide field monitor (wfm) has a 4 steradian field of view at soft x-rays to discover and localize x-ray transients and impulsive events and to monitor spectral state changes, triggering follow-up observations and providing a wealth of science in its own. loft wfm is suitable for monitoring of the outbursts and high states of cvs in the energy range of 1–50 kev. 3.1 the loft science the main loft science is the study of neutron star structure and equation of state of ultradense matter (3+ independent methods): (1) neutron star (ns) mass and radius measurements, and (2) neutron star crust properties. in addition to that, strong gravity and the mass and spin of black holes will be investigated (5 independent methods), e.g. quasi-periodic oscillation evolution, and, in the time domain, nfe line tomography and reverberation studies in bright active galactic nuclei (agns) and black hole candidates (bhcs) ((also vs nss), and relativistic precession. loft observatory science is expected to provide important observations for virtually all classes of relatively bright sources, including: x-ray bursters, high-mass x-ray binaries, xray transients (all classes), cvs, magnetars, gamma-ray bursts (grbs), nearby galaxies (smc, lmc, m31...), agn etc. in general, cvs are x-ray emitters (e.g. warner 1995). however, the properties and time evolution of their x-ray spectra strongly depend on the type and the state of activity of a given cv. although some general relations exist for a given cv type, each cv is really specific in this regard. since the proposed band of loft is 1–50 kev, those cvs with the hardest xray spectra will be the suitable targets for observing with this satellite. for example, the intermediate polars usually display the significantly larger values of kt in comparison with non-magnetic cvs (šimon et al. 2006). magnetic cv often have the hard x-ray intensity correlated with the mass accretion rate onto the white dwarf because of the accretion rate onto the magnetic poles of the wd. the dwarf nova outbursts of intermediate polars can be therefore accompanied by a strong increase of the hard x-ray intensity (e.g. do dra (szkody et al. 2002), gk per (fig. 7)). in polars (e.g. warner 1995), the hard x-ray intensity strongly increases in the optical high states. however, the case of the polar am her shows that the relation of the optical and hard x-ray emission varies for the individual episodes of the high state (šimon 2011). this emphasizes the necessity to obtain the observations of a larger ensemble of cvs to study the processes and their parameters which influence the behavior of cvs on the long timescales. we therefore expect the satellites like loft to be important instruments for investigation of magnetic cvs. figure 7: outbursts of the intermediate polar gk per in the optical band (afoev data) (a) and the hard xray band (bat/swift data (data of krimm et al. 2013). 4 conclusions both gaia and loft can contribute to study of cvs and related objects. spectrophotometry and low dispersion spectroscopy are the most important for cv analyses with the observations of gaia. both the loft wfm x-ray monitoring in the energy range of 1–50 kev and detailed lad observations with fine time resolution are important in this research field. gaia was launched on december 19, 2013. loft was not selected in competition with other esa m3 candidates, but it remains a promising candidate for m4. it is true that dense series of observations covering the intervals of several years are necessary to investigate the rising and decaying branches of outbursts and high/low states. nevertheless, a properly chosen strategy can yield suitable results of analyses even in the sampled photometric and spectroscopic data provided by gaia. we propose to obtain the statistical properties of the long-term activity of a representative ensemble of events (e.g. outbursts, high and low states) in a very large ensemble of a given type of systems. this is important for our understanding of the physical processes involved in the systems distributed in various regions of the galaxy which may have a different history of the star formation and chemical evolution. we also argue in favor of the development of more sensitive x-ray monitors that will be able to detect the long-term activity of cvs. even the available short data series obtained by the pointed x-ray observations 301 r. hudec, v. šimon of some known cvs revealed the very large variety of properties of their x-ray emission (e.g. warner 1995). the current monitors are able to observe mainly the luminous binary x-ray sources which contain the ns or the bh. since only a very few cvs are detectable by these instruments, loft can be a promising satellite to provide a sensitive x-ray monitor. acknowledgement we acknowledge partial support by ga cr grants 1339464j and 13-33324s. we also acknowledge the use of public data from swift/bat transient monitor provided by the swift/bat team. we also used the observations from the aavso international database (massachusetts, usa) and the afoev database operated in strasbourg, france. references [1] brandt, s., et al., 2013, eas publications series, volume 61, 2013, pp.617-623 doi:10.1051/eas/1361098 [2] davidson, k., et al., 1978, apj, 220, 239 [3] de boer,k., gilmore, g., hog, e., lattanzi, m.g., lindegren, l., luri, x., mignard, f., face, o., ferryman, m., de zeeuw, p.t. (eds) gaia concept and technology study report (the ’red book’) esa-sci(2000)4 pp381 (esa paris) 2000 [4] hudec, r. et al., 2013, acta polytechnica, vol. 53, no. 3, p.30 [5] eyer, l. et al., 2013, central european astrophysical bulletin, p. 115-126 [6] feroci, m. et al., 2012, experimental astronomy, volume 34, issue 2, pp.415-444 doi:10.1007/s10686-011-9237-2 [7] herbig, g. h., et al., 1965, apj, 141, 617 doi:10.1086/148149 [8] hudec, l., 2007, algorithms for spectral classification of stars, bsc. thesis, charles university, prague [9] hudec, r., šimon, v., 2007a, specific object studies for cataclysmic variables and related objects esa gaia reference code gaia-c7-tn-aio-rh-001-1. [10] hudec, r., šimon, v., 2007b, specific object studies for optical counterparts of high energy sources. esa gaia reference code gaia-c7-tn-aio-rh-002-1. doi:10.14311/ap.2013.53.0799 [11] hudec, r., šimon, v., hudec, l., 2013, acta polytechnica, vol 53, supplement, p.798 doi:10.1088/0067-0049/209/1/14 [12] krimm, h. a., et al., 2013, apjs, 209, 14 doi:10.1111/j.1745-3933.2010.00979.x [13] munari, u., et al., 2011, mnras, 410, l52 doi:10.1093/mnras/stt1340 [14] munari, u., et al., 2013, mnras, 435, 771 [15] šimon, v., mattei, j. a., 1999, a&as, 139, 75 [16] šimon, v., 2000, a&a, 360, 627 [17] šimon, v., et al., 2006, iaus, 230, 66 [18] šimon, v., hric, l., petŕık, k., et al., 2002, a&a, 393, 921 [19] šimon, v., 2010, advances in astronomy, article id. 382936 doi:10.1016/j.newast.2011.03.001 [20] šimon, v., 2011, newa, 16, 405 [21] steiner, j. e., diaz, m. p., 1998, pasp, 110, 276 [22] szkody. p., et al., 2002, aj, 123, 413 [23] van amerongen, s., van paradijs, j., 1989, a&a, 219, 195 [24] walker, m., & chincarini, g., 1968, apj, 154, 157 302 http://dx.doi.org/10.1051/eas/1361098 http://dx.doi.org/10.1007/s10686-011-9237-2 http://dx.doi.org/10.1086/148149 http://dx.doi.org/10.14311/ap.2013.53.0799 http://dx.doi.org/10.1088/0067-0049/209/1/14 http://dx.doi.org/10.1111/j.1745-3933.2010.00979.x http://dx.doi.org/10.1093/mnras/stt1340 http://dx.doi.org/10.1016/j.newast.2011.03.001 introduction the esa gaia mission astrophysics with gaia statistical properties of the long-term activity of cvs, perspectives of histograms for gaia rare flares (different from outbursts in dwarf novae) in cvs ultra-low dispersion spectra the color indices from rp/bp spectra searches for cycles or periods esa loft the loft science conclusions 322 acta polytechnica ctu proceedings 1(1): 322–325, 2014 322 doi: 10.14311/app.2014.01.0322 final remarks giulio auriemma1,2 1universiẗı¿œ degli studi della basilicata, potenza,italy 2infn sezione di roma, rome, italy corresponding author: giulio.auriemma@cern.ch this year we have to celebrate the 50th anniversary of the book “the structure of the scientific revolutions” by thomas kuhn that has been published in 1963.. in this book, the more influential science philosopher of the last century, changed the old view of the development of the natural sciences as a linear process accumulation of knowledge into a substantially discontinuous passage from one paradigm to a new one. two drastic changes of paradigm occurred near the beginning of last centuries with the shift from newton’s absolute time to relativistic space-time and from laplace determinism to quantum mechanics. around the middle of the same century another crucial shift of paradigm was the introduction of gauge symmetries, that led to the formulation of the standard model of electromagnetic, weak and strong interactions (sm for brief) [1], whose complete experimental verification has been given recently by lhc [2, 3]. the carriers of those three forces are spin-1 particles, namely the photon for the electromagnetic, the gluon for the strong and the w±,z for the weak one. it is well known that the first two particles are massless, while the others, that were discovered at lep. have large masses (∼ 80 − 90 times the mass of the proton), which can explain the weakness of the corresponding interactions. since the sm does not include gravity, there is no hint in it for the mass of the particles, that are put “by hand” in the lagrangian. what is now usually called the “higgs mechanism” [4, 5] is the possibility that the mass of particles, in particular that of the weak bosons, could be originated by the coupling with a universal scalar field, which carriers would be a weakly interacting spin-0 particles very similar to the one recently observed at lhc. nevertheless even a confirmed discovery of the higgs would not be the end of the story, because thanks to astronomy and cosmology, we have strong observational evidences for phenomena like inflation dark matter dark energy baryon asymmetry that are not understood in the framework of the minimal sm. this justify the diffuse opinion (hope?) that a lot of new physics remains to be discovered in the sky and laboratory. 1 cosmological parameters the results presented by rubiño-martin [6] on behalf of the planck collaboration is one of the highlights of this conference. the new measurement of the hubble constant from the fit of the cmbr obtained from the planck satellite 2013 data is h0 = (67.3 ± 1.2) km s−1 mpc−1 is fully compatible with the previous value of 70 ± 2.2 published by wmap collaboration. however the central value of h0 is changed by non-negligible factor (-4%) that indicates that the age of the universe is effectively t0 = (13.82 ± 0.12) gy, about 100 million longer then the previous estimate from wmap. incidentally that caused a funny “communication” problem with the media, because it was public ally announced on being the “universe older then the big bang”. figure 1: reproduced from ref. [6] the tension of the h0 value between hst key project (cepheid+sne ia) and the cmbr anisotropy fit, shown in fig. 1, that was marginal (−1.2σ) with 322 http://dx.doi.org/10.14311/app.2014.01.0322 final remarks wmap, is now at a level of −2.5σ with the new planck data. not a suspicious level yet, but close to become interesting. table 1: universe composition from cmbr before planck after planck dark energy 72.8% 68.3%(−4.5% ⇓) dark matter 22.7% 26.8%(+4.1% ⇑) ordinary matter 4.5% 4.9%(+0.4% ') total 100% 100% a similar situation is observed in the determination of the λcdm composition breakdown as shown in table 1. the uncertainty on the fractions is estimated by the planck collaboration to be about ±3% [7]. but it should be also stressed that this fraction are obtained with a multivariate fit over a constrained domain, in which correlations play an essential role. therefore the fit error given independently for each of the variables could be strongly underestimated. generally speaking a reliable error on the determination of each parameter should be obtained using the full covariance matrix of the fit, if definite positive. 2 dark energy nino panagia [8], presenting the most recent results of the hubble supernovae cosmology project (scp), strongly stressed that the λcdm model with ωλ = 0.729 ± 0.014 indicating a present universe dominated by the mysterious dark energy, are all based on the assumption that the explosive properties of the sne ia do not depend from redshift, up to z ' 1.4. both the scp and ground based telescope surveys of light curves and spectroscopic distributions of hundreds of supernovae, consistently found that the supernovae around z ≈ 0.5 appear to be ≈ 0.3 mag dimmer then expected from a flat universe with ωλ = 0. what if a cosmical conspiracy made the sne at redshift ≈ 0.5 intrinsically dimmer? as predictable this pebble thrown in the placid pool (near stockholm) started a heated discussion. many in the audience supported the argument that the good fit of distance modulus vs. redshift curve [9] and of the tt spectrum of cmb fluctuations. in the discussion following the presentation of his paper, nino questioned from a philosophical point of view that a successful fit is not to be taken as a definitive proof of validity for a theory. his point is absolutely correct, but popper would object that a scientific theory can be falsified, not affirmed. therefore we can exclude theories that do not fit the data, but not viceversa. on the same ground i think that would be very hard to fit the planck data with ωλ = 0. harmes [10] has presented an interesting paper on a possible identification of the dark energy with solitonic primordial gravitational waves in the framework of quantum gravity with one warped extradimension, similar to the russel saunders one. beside the specific model presented in that talk, that may explain both strength and time evolution observed by planck [11], it has been speculated that dark energy could be the present manifestation of long lasting relics, with lifetime much larger then the hubble time, produced in the planck era (tu . 5 × 10−44 s) by quantum gravity effects. 3 dark matter astrophysical evidence for dark matter is now overwhelming. in a recent paper bahcall and kuilier [12] tracing the mass-to-light ratio with dynamical and lensing methods, show that this ratio m/l � 1 is nearly constant for scales from 350h.1 kpc up to 22h−1 mpc, while the fraction of the stellar mass over the total mass remains of the order of few percent over all scales and environments. it worth noticing that quantitatively the observed mass-to-light ratio on large scales gives an estimate of the mass density of the universe ωm = 0.26 ± 0.02, slightly smaller (≈ 1.5σ) then the planck value [6]. after the discovery of the higgs particle, dark matter is considered to be the most compelling evidence for new physics [13, 14]. particle theory offers several type of objects that could play the role of observed dark matter, as for example axions and majorana neutrinos, but excellent theoretical candidates are the supersymmetric neutral partners of the sm particles. lhc has definitively the possibility of discovering these particles if their mass is in the tev range. however, as i said in my talk, there is a fruitful confrontation among astronomical observatories and accelerators but not a competition, because astronomy cannot demonstrate that wimp’s are supersymmetrical particles and accelerators cannot prove that a certain kind of particle really is the dark matter. several talks have described the status of the search for wimp annihilation signal in γ-rays survey. in particular morselli [15] presented on behalf of the fermilat collaboration several interesting results both on the diffuse component and of the gc region. at present we have some indications that the allowed mass scale for wimp as well as the one of the susy neutralino at accelerators is to be found above several hundreds of gevs. 4 uhe neutrinos (?) the detection of neutral particles in the km3 detector icecube [16] in antarctica with an energy release ≥ 50 323 giulio auriemma tev is something waited since a long time. the study of a deep underwater muon and neutrino detector (dumand) in the pacific ocean, off the shore of the island of hawaii, usa, started in the late ’70s [17]. but unfortunately after many years of r&d the project was interrupted in 1996, after the decision of the doe science committee of not funding its deployment. in the meanwhile many other projects were initiated in many parts of the world, and often the collaborations have presented progress reports to the vulcano conferences. the icecube collaboration has announced last april the detection of two contained showers with an energy estimated from the total number of photo-electrons & 1 pev. it is worth noticing that this energy estimate is based upon the assumption that all the light is emitted in the electromagnetic cascade. the total number of events (28) detected until now by the detector [18] seems to give the first clear indication of nongeophysical and extra-solar origin neutrinos. 1. fargion[19] has proposed in his talk some intriguing possible mechanism originating these events, that could be tested in the future, when a larger number of events will be hopefully detected. it is however to be stressed that these events are really exceptional under many aspects. if they are true neutrinos the c.m.s energy of their interaction: √ s = √ 2mneν = 316 − 1400 gev � mz is order of magnitudes larger than any neutrino-nucleon interaction observed in laboratory until now. it has been proposed in the past [20, 21, 22] that different kind of new physics could enhance substantially in this energy range. moreover the detected showers could have been originated not by neutrinos but by new physics messenger, such as for example susy particles. acknowledgments i am sure to interpret the feelings of all the participants to this very successful meeting thanking the scientific committee for having assembled such an interesting program; all the speakers for their efforts in communicating results and ideas; daniela and francesco for the precious help given to run smoothly the meeting; valentina, lisa, flavia and alessandro for the good music and readings, that gave us a moment of true ... “felicitas” the weather in mondello not making us to regret of being segregated inside the lecture room; and finally franco ... everybody knows why! references [1] g. altarelli, standard model of particle physics, in: j.-p. francoise, g. l. naber, , t. s. tsun (eds.), encyclopedia of mathematical physics, academic press, oxford, 2006, pp. 32-38. arxiv:hep-ph/ 0510281. [2] g. aad, et al., observation of a new particle in the search for the standard model higgs boson with the atlas detector at the lhc, phys.lett. b716 (2012) 1-29. arxiv:1207.7214 doi:10.1016/j.physletb.2012.08.020 [3] s. chatrchyan, et al., observation of a new boson at a mass of 125 gev with the cms experiment at the lhc, phys.lett. b716 (2012) 30–61. arxiv:1207.7235. doi:10.1016/j.physletb.2012.08.021 [4] p. w. higgs, broken symmetries, massless particles and gauge fields, phys. lett. 12 (1964) 132– 133. doi:10.1016/0031-9163(64)91136-9 [5] f. englert, r. brout, broken symmetry and the mass of gauge vector mesons, phys. rev. lett. 13 (9) (1964) 321–323. doi:10.1103/physrevlett.13.321 [6] j. rubiño-martin, planck results : a review, in: these proceedings, 2013. [7] p. ade, et al., planck 2013 results. xvi. cosmological parameters, arxiv:1303.5076. [8] n. panagia, the hubble space telescope cluster supernova survey, in: these proceedings, 2013. [9] n. suzuki, et al., the hubble space telescope cluster supernova survey. v., astroph. j. 746 (2012) 85. arxiv:1105.3470, [10] b. harmes, gravitational waves and dark energy, in: theese proceedings, 2013. [11] p. l. biermann, b. c. harms, can dark energy be gravitational waves? arxiv:1305.0498. [12] n. a. bahcall, a. kulier, tracing mass and light in the universe: where is the dark matter?, mon. note r. astron. soc. in press. arxiv:1310.0022. [13] g. altarelli, the higgs: so simple yet so unnatural, in: talk given at the nobel symposium on lhc results, krusenberg, sweden, 3-17 may,, 2013. arxiv:1308.0545. 324 http://arxiv.org/abs/hep-ph/0510281 http://arxiv.org/abs/hep-ph/0510281 http://arxiv.org/abs/1207.7214 http://dx.doi.org/10.1016/j.physletb.2012.08.020 http://arxiv.org/abs/1207.7235 http://dx.doi.org/10.1016/j.physletb.2012.08.021 http://dx.doi.org/10.1016/0031-9163(64)91136-9 http://dx.doi.org/10.1103/physrevlett.13.321 http://arxiv.org/abs/1303.5076 http://arxiv.org/abs/1105.3470 http://arxiv.org/abs/1305.0498 http://arxiv.org/abs/1310.0022 http://arxiv.org/abs/1308.0545 final remarks [14] l. bergstrom, cosmology and the dark matter frontier, in: invited talk at the nobel symposium on lhc physics, krusenberg, sweden, may 13-17,, 2013. arxiv:1309.7267. [15] a. morselli, last results from the fermi-lat gamma-ray telescope, in: these proceedings, 2013. [16] m. aartsen, et al., first observation of pev-energy neutrinos with icecube, physical review letters 111 (2) (2013) 021103. doi:10.1103/physrevlett.111.021103 [17] a. roberts, r. donaldson (eds.), 1976 dumand summer workshop, 1977. [18] m. spurio, photonic, neutrino and particle astronomy as messenger of the universe, in: these proceedings, 2013. [19] d. fargion, uhecr correlated tev anisotropy connection with pev neutrino shower events, in: this proceedings, 2013. [20] m. s. carena, d. choudhury, s. lola, c. quigg, manifestations of r-parity violation in ultrahighenergy neutrino interactions, phys.rev. d58 (1998) 095003. arxiv:hep-ph/9804380, doi:10.1103/physrevd.58.095003 [21] m. kachelriess, m. plumacher, ultrahighenergy neutrino interactions and weak scale string theories, phys.rev. d62 (2000) 103006. arxiv:astro-ph/0005309. doi:10.1103/physrevd.62.103006 [22] r. gandhi, ultrahigh-energy neutrinos: a review of theoretical and phenomenological issues, nucl.phys.proc.suppl. 91 (2001) 453–461. arxiv:hep-ph/0011176. doi:10.1016/s0920-5632(00)00975-0 hope to meet you all at the next frascati workshop 325 http://arxiv.org/abs/1309.7267 http://dx.doi.org/10.1103/physrevlett.111.021103 http://arxiv.org/abs/hep-ph/9804380 http://dx.doi.org/10.1103/physrevd.58.095003 http://arxiv.org/abs/astro-ph/0005309 http://dx.doi.org/10.1103/physrevd.62.103006 http://arxiv.org/abs/hep-ph/0011176 http://dx.doi.org/10.1016/s0920-5632(00)00975-0 cosmological parameters dark energy dark matter uhe neutrinos (?) 123 acta polytechnica ctu proceedings 2(1): 123–127, 2015 123 doi: 10.14311/app.2015.02.0123 photometric and spectroscopic investigation of the dwarf nova hs 0218+3229: a short review n. katysheva1, s. shugarov1,2, n. borisov3, m. gabdeev3, p. golysheva1 1sternberg astronomical institute of the moscow state university, 119991 moscow russia 2astronomical institute of the slovak academy of sciences, 059 60 tatranská lomnica, slovakia 3special astrophysical observatory of the russian academy of sciences, 369 167 nizhnij arkhyz, russia corresponding author: natkat2006@mail.ru abstract this paper is devoted to the study of the cataclysmic variable hs 0218+3229 using the photometric and spectroscopic observations. keywords: cataclysmic variables dwarf novae optical photometry spectroscopy individual: hs 0218+3229. 1 introduction cataclysmic variables (cvs) are highly evolved close binaries consisting of a white dwarf (wd) accreting matter from a red dwarf companion. this matter creates an accretion disc around the white dwarf. dwarf novae (dne) are a subclass of cataclysmic variables, whose members experience outbursts either caused by a thermal-viscous instability in the accretion disc (di) or by a sudden increase in mass transfer rate. outbursts of dne last a few days. in gcvs (samus et al., 2007-2012) dne are described as stars of the type ugss (u gem–ss cyg), with the recurrent time between outbursts lasting from days to several years. hs 0218+3229 (α2000 = 02 h21m34s; δ2000 = +32 ◦ 43’ 24”) was discovered spectroscopically during the large-scale search for cvs in the hamburg quasar survey by gänsicke et al. (2002). rodriguez-gil et al. (2009) carried out photometric and spectroscopic observations of hs 0218+3229 from 2000 to 2005 and determined its orbital period to be 0.d297229661(1). the phase light curve (lc) in the r passband exhibits ellipsoidal modulation. the k5v secondary component gives a contribution of 80-85% to the r light. an analysis of the time-resolved optical spectroscopy and r passband photometry provided some parameters of hs 0218+3229: a mass ratio of 0.52< q <0.65, a white dwarf mass of 0.44< m1/m� <0.65, a secondary component mass of 0.23< m2/m� <0.44, the orbital inclination i = 59o ± 3o. the distance to the system was estimated to be 0.87–1.0 kpc. hs 0218+3229 has been identified as the x-ray source 1rxs j022133.6+324343 and the 2mass source j02213348+3243239. 2 photometric observations of hs 0218+3229 2.1 optical outbursts in 2006 the object was discovered independently by s. antipin using the photo plates of the photographic archive at the sternberg astronomical institute (sai) and identified as a dn. one outburst with an amplitude of about 4.5 mag in the bph passband was detected in september 1980. another outburst was detected in 2002 by wills from neat database (2002). golysheva et al. (2012, 2013) described photometric and spectroscopic observations of hs 0218+3229 from 2006 till 2010. ccd photometry in ubvrcic passbands was carried out at the crimean laboratory of sai (nauchny) and at the stará lesná observatory of the astronomical institute of the slovak academy of sciences. the magnitudes of comparison stars were determined in golysheva et al. (2012). in october 2007 they detected the outburst of hs 0218+3229 with an amplitude of 4 mag and duration of 15–16 days (jd 2454380–395). the shape of the outburst lc was more symmetric than for other dne. its asymmetry coefficient (the ratio of the duration of the ascending branch to the descending branch) was ∼0.22. a new ephemeris was calculated using the initial epoch of the minimum from rodriguezgil et al. (2009). the epoch corresponds to the inferior conjunction (the secondary is in front of the wd): hjdmin = 2453653.0286 + 0.2973559 ×e (1) the orbital period of the object (7.13 h) largely ex123 http://dx.doi.org/10.14311/app.2015.02.0123 n. katysheva et al. ceeds the average orbital periods of dne, which are in the range 1.5–4 hours. the maximum v magnitude of the 2007 outburst was determined as vmax=12.34±0.05 mag. the maximum magnitudes in other passbands were estimated by extrapolation as u=11.3±0.2 mag, b=12.3±0.2 mag, rj 'ij =12.2±0.2 mag. during 2011–13 we continued observations of the object. as seen from bvrcic light curves (lcs) of the object, obtained in 2006-13 and presented in fig. 1, a new outburst occurred in 2013. its v-maximum magnitude was the same as in 2007 outburst. a sudden 0.5 mag increase of brightness in the b passband was detected on january 31, 2011. the bvrcic plots in fig. 1 show long-term brightness variations during quiescence with a characteristic timescale of a few tens of days and an amplitude of ∼0.2 mag, noted in golysheva et al. (2012). figure 1: the bvrcic lcs of hs 0218+3229 in 2006–2013. the aavso data are marked by grey circles. our observations of the last outburst of hs 0218+3229, detected on september 5, 2013 (vsnet, 2013), were obtained at the stará lesná observatory. taking into account that the object exhibited outbursts also in 2002 and 2007, it is possible to estimate the recurrent time between outburst to be 5–6 years. fig. 2 shows the lcs of the outbursts in 2007 (upper panel) and 2013 (lower panel). the nightly lcs are inserted in separate panels. left lc, obtained directly before the outburst in 2007, shows half of orbital wave with an ellipsoidal effect. this effect is not visible on the other nightly lcs because of the presence of the outbursting disk. in the framework of the di model, the shape of the outburst lc of hs 0218+3229 can be explained by a low mass transfer rate in the system and ”inside-out” outburst in accordance with smak (1984). outburst starts from the inner parts of the disk and extends outward. the 2013 outburst shape is similar to the previous one in 2007. the nightly bv lcs are shown in fig. 2. figure 2: the outburst lcs in 2007 (top) and 2013 (bottom) in v (black points and circles) and b (grey points). the aavso v data are marked by grey circles. the ∆ubv rcic outburst lcs, near the maximum brightness, are presented in fig. 3. all lcs look similar. figure 3: the lcs after removal of the declining trend in brightness during the outburst, on september, 8, 2013. the lcs are displaced by arbitrary magnitudes. 124 photometric and spectroscopic investigation of the dwarf nova hs 0218+3229: a short review the quiescence lcs in the bvrcic passbands, folded with the ephemeris (1), are shown in fig. 4. double wave is clearly seen in the vri passbands, and less clearly in the b passband. it should be noted that the heights of the maxima (in phases 0.25 and 0.75) and depth of the minima (in phases 0 and 0.5) can change from night to night. the displacement of the second minimum around phase 0.5 is clearly seen. this feature was noted also by rodriguez-gil et al. (2009). golysheva et al. (2012, 2013) noted a large uvexcess in quiescence in 2010. during the 2007 outburst the position of the star in the two-colour diagram was close to that of an absolute black body with a temperature of about 15000 k. at the minimum brightness the position of the variable corresponded to a star of spectral class k5v-k6v. more detailed information about two-colour diagrams can be found in the paper of golysheva et al. (2012). figure 4: the average orbital phase bvrcic lcs in 2006–2012, folded with ephemeris (1). individual observations are marked by grey points, mean values by black points. 2.2 the analysis of the aavso data we also present data by the american association of variable star observers (aavso), whose observers were asked to observe the object to support the hubble observations (hst/cos, the aavso ”alert 471” (2012) about monitoring of hst targets). the hst/cos observations were obtained in december 2012. the main bulk of aavso observations were carried out in a clear filter. because of a very good time coverage, the orbital lcs were very well covered by observations in some nights from september 17, 2012 till january 2, 2013. we reduced the data to heliocentric values. the nightly lcs in september and october 2012 are presented in fig. 5 and folded lcs are shown in fig. 6. in spite of the error 0.02–0.03 mag, the shape of the nightly lcs is distinctive. the depth of the primary minimum (marked by 1) is smaller than the depth of the secondary minimum (marked by 2) and the hump after the primary minimum is brighter than the hump after the secondary minimum. this phenomenon is especially evident in the orbital folded phase lcs (fig. 6). figure 5: the aavso lcs obtained in september and october 2012. the primary and secondary minimum are marked by 1 and 2, respectively. figure 6: the average lcs in september–october 2012, folded with ephemeris (1). the notation is the same as in fig. 4. 125 n. katysheva et al. 3 spectroscopic observations our spectroscopic observations of hs 0218+3229 was taken in 2010 and 2012, when the object was in quiescence. golysheva et al. (2013) obtained two spectra of hs 0218+3229 on september 17/18, 2010 in the prime focus of the russian 6-m telescope bta of the special astrophysical observatory of the russian academy of sciences (sao ras) at the spectroscopic mode of the scorpio multi-mode focal reducer (afanasiev, moiseev, 2005) with the long slit and ccd-camera eev ccd 42-40 (2048×2048 pixels) with the exposures of 300 sec. our new spectroscopic observations of the objects were carried out on november 5, 2012 with the same reducer at the 6-m bta telescope. the vphg1200g grating (1200 grooves/mm) were used in both cases. the spectroscopic resolution ∆λ = 5.0 å in the wavelength interval 3950–5700 å was achieved. the spectra were reduced with sky and bias substraction and division by a flat-field frame. the discovery spectrum of the object obtained in 2000 (rodriguez-gil et al., 2009) revealed the absorption-line spectrum of the k5 v secondary, accompanying by the balmer and he i emission lines, arising in the accretion disk of the white dwarf primary. the 2010 spectrum was different. emission line of neutral helium λ5875 å became stronger, emission balmer decrement more flat and the features of the red companion became less pronounced (golysheva et al., 2012, 2013). these changes were probably caused by enhancing of a mass transfer rate and veiling of the spectrum of the secondary component. our spectrum obtained on november 5, 2012 is similar to the spectrum taken in 2000. the red star manifested itself again. 4 discussion and conclusions four outbursts of hs 0218+3229 in 1980, 2002, 2007, 2013 were detected up to now. the 2002 outburst (wills, 2002) was found in the database neat. the duration of 2007 outburst was 15–16 days and its asymmetry coefficient ∼0.22. the maximal amplitudes were 5 mag in the u, b passbands and 4 mag, 3.5 mag, 3 mag in the v, r, i passbands, respectively. it is possible to estimate the recurrence time of the outbursts to 5–6 years. we determined the orbital period to be 0.d2973559(10) and updated the ephemeris of the object. two-colour diagrams, presented in golysheva et al. (2012, 2013), showed that the object was bluer during its outburst than in quiescence. it was caused by cthe additional ontribution of a bright accretion disk to the total luminosity. outburst colour temperature was about 15 000 k. colour indices in quiescence correspond to a star k5 v. as a result, this cataclysmic variable was classified as a very rare subtype of ugsstype dn with a low accretion rate and sparse and more symmetric (”inside-out”) outbursts than it is usual in u gem-type and ss cyg-type of dne. figure 7: the spectrum of hs 0218+3229 on november 5, 2012. acknowledgement we thank to the large telescopes program committee of the sao ras for supporting our program of spectroscopic study of cvs. we would like to express our deep gratitude to dr. d. chochol for his valuable suggestions and critical reading of the manuscript. we are grateful to many amateur astronomers whose observations from the aavso international database were used in this research. nk thank to the soc for possibility to make a report. the work was supported partially by grants: nsh-1675.2014.2, rfbf-11-11-02-00258, 12-0200186, 12-02-97006 and vega grant 2/0002/13. references [1] samus, n., durlevich, o.v., kazarovets, e v. et al.: 2007-2012, general catalogue of variable stars, vizier on-line data catalog: b/gcvs. [2] gänsicke, b. et al.: 2002, asp conf. ser., 261, 190. [3] rodriguez-gil, p., torres, m.a.p., gänsicke, b.t. et al.: 2009, a&a, 496, 805. [4] wills, p.: 2002, near-earth asteroid tracking: http://neat.jpl.nasa.gov. [5] vsnet-alert.: 2013, http://ooruri.kusastro. kyoto-u.ac.jp/mailarchive/vsnet-alert/16355. 126 photometric and spectroscopic investigation of the dwarf nova hs 0218+3229: a short review [6] golysheva, p.yu., antipin, s.v., zharova, a.v. et al.: 2012, astrophysics, 55, 208. doi:10.1007/s10511-012-9229-6 [7] golysheva, p., katysheva, n., shugarov, s. et al.: 2013, ceab, 37, 345. [8] aavso-alert.: 2012, http://www.aavso.org/ aavso-alert-notice-471 [9] afanasiev, v., moiseev, a.: 2005, astron. lett., 31, 194. doi:10.1134/1.1883351 [10] smak, j.: 1984 pasp, 96, 575. doi:10.1086/131295 127 http://dx.doi.org/10.1007/s10511-012-9229-6 http://dx.doi.org/10.1134/1.1883351 http://dx.doi.org/10.1086/131295 introduction photometric observations of hs 0218+3229 optical outbursts the analysis of the aavso data spectroscopic observations discussion and conclusions 71 acta polytechnica ctu proceedings 2(1): 71–75, 2015 71 doi: 10.14311/app.2015.02.0071 photometry and multipolar magnetic field modeling of polars: by camelopardalis and fl ceti p. a. mason1,2, a. g. zhilkin3, d. v. bisikalo3, s. gomez1, j. morales1, e. l. robinson4 1department of physics, university of texas at el paso, el paso tx, usa 2department of mathematics and physical sciences, new mexico state university dacc, las cruces nm, usa 3institute of astronomy, russian academy of sciences, moscow, russia 4department of astronomy, university of texas at austin, austin tx, usa corresponding author: pmason@nmsu.edu abstract we present new broad band optical photometry of two magnetic cataclysmic variable stars, the asynchronous polar by camelopardalis and the short period polar fl ceti. observations were obtained at the 2.1-m otto struve telescope of mcdonald observatory with 3s and 1s integration times respectively. in an attempt to understand the observed complex changes in accretion flow geometry observed in by cam, we performed full 3d mhd simulations assuming a variety of white dwarf magnetic field structures. we investigate fields with increasing complexity including both aligned and non-aligned dipole plus quadrupole field components. we compare model predictions with photometry at various phases of the beat cycle and find that synthetic light curves derived from a multipolar field structure are broadly consistent with optical photometry. fl ceti is observed to have two very small accretion regions at the foot-points of the white dwarf’s magnetic field. both accretion regions are visible at the same time in the high state and are about 100 degrees apart. mhd modeling using a dipole plus quadrupole field structure yields quite similar accretion regions as those observed in fl ceti. we conclude that accretion flows calculated from mhd modeling of multi-polar magnetic fields produce synthetic light curves consistent with photometry of these magnetic cataclysmic variables. keywords: cataclysmic variables magnetic fields polars optical photometry individual: by cam individual: fl cet. 1 introduction we report on new observations of magnetic cataclysmic variable stars (mcvs). photometric observations of two mcvs: by cam and fl ceti, were obtained at mcdonald observatory using the otto struve 2.1-m telescope. the data consists of 1 or 3 second exposures and no dead time. a broad band filter covering the visual spectrum, bvr, was used in order to obtain the highest possible time resolution. we describe progress towards understanding the structure of accretion flows in mcvs, obtained through the use of full 3d mhd simulations assuming a variety of magnetic field structures. 2 by camelopardalis by cam contains a highly magnetic white dwarf which rotates slightly (about 1 %) faster than its companion (donor star) orbits (mason et al. 1989, 1998; silber et al. 1992, 1997). hence, it is a member of the class of mcvs known as asynchronous polars or by cam stars. the binary period is 201.3 minutes while the white dwarf spin period is 199.3 minutes (mason et al. 1995). a precise orbital ephemeris was derived by schwarz et al. (2005). the vast majority of polars are synchronously locked due to the strong magnetic field of the white dwarf primary. despite the fact that by cam has a magnetic field typical of polars, it currently is in an unlocked state. mason and chanmugam (1992) and pavlenko et al. (2012) provide evidence for the synchronization of the orbit and spin periods in 1600 and 250 years respectively. by cam is well suited for the study of magnetic field structure in mcvs because we are able to take advantage of asynchronous rotation of the white dwarf with respect to the binary. synchronized polars, by definition, only accrete from one azimuthal angle. in by cam, accretion flow occurs at all azimuthal angles as the field structure moves with respect to the orbital plane over the course of a 14 day beat cycle (mason et al. 1995, 1998), see figure 1. other extensive photometric studies of by cam were conducted by andronov et al. (2008) and babina et al. (2010). 71 http://dx.doi.org/10.14311/app.2015.02.0071 p. a. mason et al. figure 1: by cam photometry. top: b-band from 7 february 2013. bottom: r-band from 12 february 2013. cycle to cycle variations are clearly evident. 2.1 by cam: mhd modeling a complex magnetic field model was proposed to explain optical photometry and polarimetry of by cam by mason et al. (1998). full 3d mhd simulations of flow structure in asynchronous polars were carried out using the russian supercomputer facility (zhilkin et al. 2012). 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 in te n si ty phase beat phase = 0.3i=30 o i=60o i=90o i=120o i=150o figure 2: synthetic light curves derived from a nonaligned dipole plus quadrupole field model at beat phase 0.3 are shown as an example. these light curves resemble some of the observed light curves of by cam, see figure 1. calculations of accretion flow patterns at 10 beat phases were performed. for pure dipole and dipole plus quadrupole field components of various dipole/quadrupole field strength ratios. figures 2 and 3 show synthetic light curves of a by cam model at two beat phases. 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 in te n si ty phase beat phase = 0.8i=30 o i=60o i=90o i=120o i=150o figure 3: same as figure 2, except for beat phase 0.8. these light curves resemble some of the observed light curves of by cam, see figure 1. figure 4 shows the magnetic induction at the surface of the white dwarf for a misaligned dipole plus quadrupole magnetic field model, currently the best fitting model for by cam. figure 5 shows the location of accretion regions for an aligned dipole plus quadrupole field structure at a particular beat phase. b/ba 0 0.5 1 1.5 2 φ/π 0 0.2 0.4 0.6 0.8 1 θ/ π 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 figure 4: the magnetic induction over the surface of the model white dwarf in by cam. dark areas represent locations of potential accretion. there are three regions, two near the magnetic equator and one at the north magnetic pole. this is a non-aligned dipole plus quadrupole field model. figure 5: an example model of by cam with dipole plus quadrupole components with accretion regions, just below the magnetic equator (see right panel) and the magnetic north pole. 72 photometry and multipolar magnetic field modeling of polars: by camelopardalis and fl ceti 3 fl ceti the original optical observations of this synchronized polar revealed deep eclipses on a period of 87 min (szkody et al. 2002). it has been observed undergoing ∼ 6 mag. eclipses in the high state, with ∼ 1.5 mag. eclipses in the low state. it has been the subject of numerous additional investigations (dubkova et al., 2003; wiehahn et al. 2004, woudt et al., 2004, schmidt et al., 2005, o’donoghue et al. 2006). katysheva and shugarov (2012) provide a concise review of past work and present additional data. we obtained optical photometry of fl ceti, during both high and low states, using the mcdonald observatory otto struve 2.1-m telescope. fast photometry allows us to resolve detailed eclipse structure resulting from the very small accretion regions on the surface of the white dwarf. eclipse’s of the accretion spots have ingress and egress times of 2-3 seconds, indicating a very small structure. polars are known for exhibiting high and low brightness states, corresponding to high and low accretion rates respectively. in some cases, low states may be periods of near zero accretion. broad-band (bvr) photometry was collected in august and september 2011 during which time fl ceti was in the high state with two active accretion regions (see figure 6). low state light curves were obtained on six nights in october 2013. v-band photometry obtained 7 october 2013 is shown in figure 7. the light is dominated by the white dwarf at a relatively constant brightness. the low amplitude variability in the out of eclipse light curve may be due to ellipsoidal variations of the secondary. that is, the visible surface area of the non-spherical donor star changes over the course of the orbital period. during the eclipse the back side of the secondary provides the only contribution to the light during both high and low states. figure 6: broad-band (bvr) photometry obtained during a high accretion state of fl ceti. outside of eclipse the light curve is dominated by cyclotron emission. during the eclipse the back side of the secondary and the disappearance of the accretion stream provide the only contribution to the optical light. in order to compare the high and low states a close up view of eclipses of the white dwarf by the secondary star are shown in figure 8, for the high state (top) and the low state (bottom). the high state eclipse involves two sharp drops during ingress and a stand still in between. the egress is symmetric about mid-eclipse of the two spots. this is interpreted as the eclipse of two small accretion spots on the surface of the white dwarf (o’donoghue, et al. 2006). one spot is eclipsed for a somewhat longer period of time than the other. low state flux is flat during the deep eclipse while the high state light curve shows a gradual decline. an explanation for the gradual fading during eclipse has been given by o’donoghue et al., (2006), that light from the accretion stream is being gradually blocked during the eclipse. figure 7: v-band photometry obtained during a low accretion state of fl ceti. figure 8: eclipses of the white dwarf by the secondary star are shown for the high state (top red points) and low state (bottom blue points). the orbital phase is shown, with phase 0 being the time of mid-eclipse. 3.1 fl ceti: mhd modeling as described in the previous section, zhilkin et al. (2012) performed calculations of mhd accretion flow in an asynchronous polar at 10 beat phases. the same calculations are used to model the field structure synchronous polars. conveniently, fl ceti has a measured magnetic field strength very similar to by cam. a first 73 p. a. mason et al. attempt at an accretion flow model resembling fl ceti is shown in figure 9. in this case, for mhd calculations, a strong quadrupole field component is assumed to be aligned with a dipole component. the donor star is at the right edge of the image and the stream phasing is similar to the pre eclipse dip observed in the light curve of fl ceti. plasma flows onto two small regions located about 100 degrees apart on the surface of the white dwarf. similar features occur for calculations involving non-aligned dipole plus quadrupole magnetic field components. figure 9: model accretion flow in fl ceti. the magnetic field structure is calculated for a centered dipole plus strong quadrupole, where the component axes are aligned. the donor star is off the far right edge on the horizontal axis. the densest part of the flow is shown in brown. accretion stream phasing is consistent with the observed pattern in fl ceti, specifically the pre-eclipse dip and the existence of magnetic poles separated by ∼ 100 degrees. in the model, one spot is at the magnetic north pole while the other is near the intersection of the magnetic equator and the orbital plane. similar flow patterns result for misaligned dipole and quadrupole components as well. 4 conclusions it is clear that pure dipolar fields are inconsistent with observations of both by cam and fl ceti. by modeling accretion flows in these binaries with higher order complex fields we are able to explain the evolution of the light curves of by cam and the location of accretion spots in fl ceti. actual fields may consist of many multipolar components. however, simple assumptions about complex field components yield results that are quite consistent with observations. acknowledgement special thanks go to franco giovannelli for the invitation to present this work in palermo and to the referee, klaus reinsch, for suggestions that improved the manuscript. we thank david buckley for discussions concerning salt data on fl ceti. this work is supported, in part, by nsf/paare grant no. 0958783. ag and db were supported by rfbr grants. references [1] andronov, i. l., et al., 2008, cejph, 6, 385 [2] babina, ju., andreev, m., pavlenko, e., 2010, aipc, 1273, 313 [3] dubkova, d. n.; kudryavtseva, n. a.; hirv, a., 2003, ibvs, 5389, 1 [4] katysheva, n., shugarov, s. 2012, memorie della societa astronomica italiana, 83, 670 [5] mason, p. a., liebert, j., and schmidt, g. d., 1989, apj, 346, 941 [6] mason p. a. and chanmugam g, 1992, aspc, 29, 216 [7] mason, p. a., et al., 1995, aspc, 85, 496 [8] mason, p. a., et al. 1998, mnras, 295, 511 [9] o’donoghue, d., et al., 2006, mnras, 372, 151 doi:10.1111/j.1365-2966.2006.10834.x [10] pavlenko, e., et al, 2013, aspc, 469, 343 [11] schmidt, g. d., et al., 2005, apj, 620, 422 doi:10.1086/426807 [12] schwarz, r., schwope, a. d., staude, a., remillard, r. a., 2005, a & a, 444, 213 [13] silber, a. d., et al., 1992, apj 389, 704 doi:10.1086/171243 [14] silber, a. d., et al., 1997, mnras, 290, 25 doi:10.1093/mnras/290.1.25 [15] szkody, p. et al. 2002, aj, 123, 430 [16] wiehahn, m., potter s. b., warner, b., woudt, p. a., 2004, mnras, 355, 689 doi:10.1111/j.1365-2966.2004.08346.x [17] woudt, p. a., warner, b., pretorius, m. l., 2004, mnras, 351, 1015 doi:10.1111/j.1365-2966.2004.07843.x [18] zhilkin a, et al.: 2012, astron. rep. 52, 4, 318 74 http://dx.doi.org/10.1111/j.1365-2966.2006.10834.x http://dx.doi.org/10.1086/426807 http://dx.doi.org/10.1086/171243 http://dx.doi.org/10.1093/mnras/290.1.25 http://dx.doi.org/10.1111/j.1365-2966.2004.08346.x http://dx.doi.org/10.1111/j.1365-2966.2004.07843.x photometry and multipolar magnetic field modeling of polars: by camelopardalis and fl ceti discussion paula szkody: how do the new light curves of by cam compare to those obtained in the past, by you and silber et al.? paul mason: the new light curves resemble the light curves from 25 years ago. however, it should be noted that recently e. pavlenko and co-workers suggest that by cam is in the process of synchronizing in about 250 years, a lot faster than the 1600 years estimated by mason and chanmugam (1992). 75 introduction by camelopardalis by cam: mhd modeling fl ceti fl ceti: mhd modeling conclusions 222 acta polytechnica ctu proceedings 1(1): 222–226, 2014 222 doi: 10.14311/app.2014.01.0222 x-ray and near-infrared spectroscopy of dim x-ray point sources constituting the galactic ridge x-ray emission kumiko morihana1, masahiro tsujimoto2, ken ebisawa2 1nishiharima astronomical observatory, center for astronomy, university of hyogo, 407-2 nishigaichi, sayo-cho, sayogun, hyogo, 679-5313, japan 2japan astrospace exporation agency, institute of space and astronautical science, 3-1-1 yoshino-dai, chuo-ku, sagamihara, kanagawa 252-5210, japan corresponding author: morihana@nhao.jp abstract we present the results of x-ray and near-infrared observations of the galactic ridge x-ray emission (grxe). we extracted 2,002 x-ray point sources in the chandra bulge field (l =0◦.113, b = 1◦.424) down to ∼10−14.8 ergs cm−2 s−1 in 2–8 kev band with the longest observation (∼900 ks) of the grxe. based on x-ray brightness and hardness, we classified the x-ray point sources into three groups: a (hard), b (soft and broad spectrum), and c (soft and peaked spectrum). in order to know populations of the x-ray point sources, we carried out nir imaging and spectroscopy observation. we identified ∼11% of x-ray point sources with nir and extracted nir spectra for some of them. based on x-ray and nir properties, we concluded that non-thermal sources in the group a are mostly active galactic nuclei and the thermal sources are mostly white dwarf binaries such as cataclysmic variables (cvs) and pre-cvs. we concluded that the group b and c sources are x-ray active stars in flare and quiescence, respectively. keywords: galaxy: bulge galaxy: disk ir x-rays. 1 introduction since the dawn of the x-ray astronomy, an apparently diffuse emission of low surface brightness has been known to exist along the galactic plane (gp; |l| <45◦, |b| <1◦), which is referred to as the galactic ridge xray emission (grxe; e.g.,worral et al. 1982; warwick et al., 1985). the x-ray spectrum is characterized by hard continuum with a strong 6.7 kev fe k emission line (koyama et al., 1986a). the origin of the grxe had been a mystery for a long time. a long-standing debate had been whether it is a truly diffuse plasma (ebisawa et al., 2001, 2005) or a sum of unresolved xray point sources (revnivtsev et al. 2006). recently, revnivtsev et al (2009) showed that ∼80% of the fe k emission line was resolved into dim point sources using the deepest x-ray observations (∼900ks) made with the chandra at a slightly off-plane region of (l, b)=(0◦.113, –1◦.424) in the galactic bulge (chandra bulge field; hereafter, cbf). if the grxe is composed of the dim x-ray point sources, new questions arise. what are the populations of the dim x-ray point sources? which class of sources contribute to the fe k emission line? we do not know the population of majority of the dim point sources due to a limited number of x-ray photons. thus, we focus on near-infrared (nir), which has almost the same penetrating power as x-rays into deep interstellar extinction toward the gp. we studied the population constituting the grxe combining x-ray data with nir data in this paper. in particular, we focus on the population contribute to the fe k line of the grxe spectrum. 2 analysis and results 2.1 x-ray 2.1.1 observation and source extraction we retrieved 10 archived data of the cbf taken with the advanced ccd imaging spectrometer (acis)-i array on board chandra with a total exposure time of ∼900 ks. we merged 10 data set and extracted 2,002 valid point sources down to ∼10−14.8 ergs cm−2 s−1 in 2–8 kev (figure 1). for all the sources, we extracted source and background events. 2.1.2 spectral fittings for the bright sources (source counts>100), we constructed the background-subtracted spectra and generated instrumental response files. we carried out spectra fittings with thermal (apec; smith et al. 2001) and 222 http://dx.doi.org/10.14311/app.2014.01.0222 x-ray and near-infrared spectroscopy of dim x-ray point sources... non-thermal models (power-law) for the bright sources. as the results, 11 bright sources with more than 1000 counts have hard power-law like hard spectra with the photon index γ∼1.5. figure 1: smoothed and exposure-corrected x-ray image of the cbf (0.5–8 kev). the field of view of the sirius (nir) observations are shown by red squares. the white circle shows the region, the result of which was published in revnivtsev et al (2009). 2.1.3 grouping 0.5 1.0 1.5 2.0 2.5 3.0 -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.2 0.4 1 2 3 4 5 q 2 = 3 *q 2 5 /q 7 5 q1=log10 q50/(1-q50) median energy (kev) c a b figure 2: x-ray color-color diagram of all detected xray sources. the converted median energy is shown in the upper x-axis. color difference shows groups defined in § 2.1.3. we constructed an x-ray color-color diagram (hong et al., 2004) using the quantiles (e25, e50, and e75) characterizing the spectral shape of each source (figure 2). here, ex (kev) is the energy below which x% of photons reside in the energy-sorted event list. e50 is equivalent to the median energy (the details are in morihana et al., 2013). here, the q1 value indicates the degree of photon spectrum being biased toward the higher (q1 > 0) or lower (q1 < 0) energy end (hard or soft spectra), and the q2 value indicates the degree of photon spectrum being less (q2 > 1) or more (q2 < 1) concentrated around the peak (broad or narrow spectra). based on the x-ray color-color diagram, we classified all the point sources into three groups, which are the group a (hard), b (soft and brooded spectrum), and the group c (soft and peaked spectrum). 2.1.4 global spectral fittings furthermore, we performed global spectral fittings in 0.5–8 kev for the composite spectrum for each group in a similar manner in § 2.1.2 (figure 4). for the group a spectrum, neither a power-law nor a thin-thermal plasma model reproduced the spectrum well, respectively because of the excess emission at 6.7 kev or a flatness of the continuum. so we fitted the spectrum with a combination of the two models, which was successful. the equivalent width (ew) of the fe k feature becomes larger as the flux decreases (figure3), which suggest that the thermal component becomes strong against the non-thermal component as the flux decreases. for the group b spectrum, several emission lines are seen, including the 6.7 kev emission line from fexxv and 2.5 kev from sxv. this set of emission lines indicates a multiple-temperature plasma, and indeed the spectrum was reproduced well with two thin-thermal plasma components, but not with one component. for the group c spectrum, several emission lines are also seen. unlike the group b sources, the fexxv emission at 6.7 kev is absent. we fitted the spectrum using the same model with group b. 0 200 400 e q u iv a le n t w id th ( e v ) 10-1610-1510-1410-13 fx (ergs cm -2 s-1) figure 3: equivalent width of the fe k line against the decreasing flux in 2-8 kev above which the cumulative combined spectra in the group a were constructed. the 1σ statistical uncertainty is shown for each data. 2.2 nir imaging 2.2.1 observation and source extraction to identify the x-ray point sources with nir, we carried out nir observations using simultaneous infrared for unbiased survey (sirius; nagayama et al., 1999) on the infrared-survey facility (irsf) 1.4 m telescope 223 kumiko morihana, masahiro tsujimoto, ken ebisawa c o u n ts s -1 k e v -1 0 .0 1 1 0 -3 1 0 -4 -2 0 2 χ energy (kev) 0.5 1.0 2.0 3.0 5.0 (a) a c o u n ts s -1 k e v -1 0 .0 1 1 0 -3 1 0 -4 -2 0 2 χ energy (kev) 0.5 1.0 2.0 3.0 5.0 (b) b c o u n ts s -1 k e v -1 0 .0 1 1 0 -3 1 0 -4 -2 0 2 χ energy (kev) 0.5 1.0 2.0 3.0 5.0 (c) c figure 4: composite spectra and the best-fit global model of the group (a) a, (b) b, and (c) c. the lower panel shows the residuals of the data to the fit. the best-fit parameters can be found in table 1. table 1: best-fit parameters for global spectral model in 0.5–8.0 kev group nh (1)1 kbt (1)2 nh (2) kbt (2) z3 γ4 χ2/d.o.f. (1022 cm−2) (kev) (1022 cm−2) (kev) a 1.09+0.39−0.50 6.65 +3.24 −3.03 2.46 +2.35 −0.58 ... 0.97 +0.36 −0.32 1.29 +0.18 −0.40 205.36/504 b 0.75+0.06−0.05 0.74 +0.54 −0.45 0.80 +0.22 −0.18 7.87 +1.86 −4.84 0.99 +0.33 −0.29 ... 98.68/103 c 0.70+0.18−0.11 0.78 +0.04 −0.03 0.04 +0.05 −0.04 4.50 +0.65 −0.35 0.15 +0.16 −0.12 ... 85.86/102 1interstellar extinction column density for the first component (the lower temperature component for the twotemperature model. 2plasma temperature for the second component (the higher temperature component for the two-temperature model). 3metal abundance relative to the solar value for the thermal component. 4photon index for the power-law model as the second component. in south africa astronomical observatory. we covered the cbf as we show in figure 1. we extracted nir source with a 3σ level using sextractor version 2.8.6, which are 52312 (j), 61,188 (h), and 65,051 (ks) sources down to ks∼16 mag. for asymmetry and photometry correction, we rendered the two micron all sky survey (2mass) point source catalog. for asymmetry, the sirius positions are determined at an accuracy of the pixel size of sirius (0.45′′). 2.2.2 cross correlation we search for possible nir counterparts for all the xray point sources using 2mass and the sirius. we searched nir counterpart sources within 1σ error circle (x-ray-2mass source; 1.3′′, x-ray-sirius source; 1.2′′). when there are two or more sources within the 1σ circle, we assumed the closest one to be the counterpart. then, we finally recognized 222 x-ray sources to have nir counterpart within 1σ circle (∼11% of all the x-ray point sources). 2.3 nir spectroscopy we conducted nir spectroscopy observation for some selected objects in the cbf using subaru/multi-object infrared camera and spectrograph (moircs). we selected 51 sources for spectroscopy in ks-band based on x-ray hardness and source variability. figure 5 shows examples of nir spectrum. we finally obtained 33 nir spectra of the x-ray point source in the cbf. combined x-ray results, there are two type of sources in the cbf, which are (1) sources with hi (brγ) and co absorption features in nir spectra and hard x-ray spectra, (2) sources only with co absorption features in nir spectra and soft x-ray spectra. from these properties, type (1) sources are k or m spectral type stars and type (2) sources are m spectral type stars. 21000 22000 23000 24000 wavelength (angstroms) b a b a c c coco co cohi hei cainai n o rm a li z e d i n te n s it y figure 5: examples of ks-band spectra in the cbf. upper label of each spectrum shows groups in § 2.1.3 defined by x-ray properties. 224 x-ray and near-infrared spectroscopy of dim x-ray point sources... 3 discussions & conclusion we now discuss likely populations in each group based on the results presented above. the transition from one class to another is continuous along the color and flux, so the groups are naturally a mixture of sources of different classes. first, we consider that the group a sources are mostly mixture of active galactic nuclei (agns) and white dwarfs (wds) binaries, each responsible for the power-law and thermal plasma components in the composite spectrum (figure 4). in the spectral fitting, the power-law component has a spectral index of 1.29+0.18−0.40, which is similar to the typical spectral form of agns (table 1 e.g.; rosati et al., 2002). for the thermal component of the composite spectrum of the group a, a strong fe k feature and 6.7 kev plasma temperature are seen. both of these features are observational characteristics of magnetic cvs (ezuka & ishida 1999). other classes of wd binaries, such as dwarf novae, pre-cvsmay also considered for the likely classes. precvs are poorly recognized class of sources, which are detached binaries of a wd and a late-type star, unlike conventional cvs that are semi-detached systems. some of them show strong fe k emission in the hard x-rays (matranga et al. 2012). in fact, showed nearinfrared spectra of selected x-ray sources presented in this paper, in which some thermal a sources do not exhibit the brγ emission that is typical for the conventional cvs (dhillon et al. 1997). we thus consider that pre-cvs, most of which do not show the brγ emission (howell et al. 2010; schmidt& mikoajewska 2003), also account for at least some fraction of the thermal source population in the group a. the group b and c sources are galactic sources with a soft thermal spectrum, and we consider that most of them are likely to be x-ray active stars. the composite spectra of these two groups were fitted with two plasma components. in these groups, the spectra type of the sources are kor m-type spectra, which has no brγ emission line. we speculate that the differences between b and c is that most c sources represent active binary stars in the quiescence, while most b sources represent those during flares. from these things, we consider that the group b and c sources are x-ray active stars in flare and quiescence, respectively. acknowledgement this research has made use of public data obtained from chandrax-ray center which is operated for nasa by the smithsonian astrophysical observatory. this work has also made use of software from high energy as astrophysics science archive research center (heasac) which is provided by nasa goddard space flight center. references [1] dhillon, v. s., marsh, t. r., duck, s. r., & rosen, s. r. 1997, minareas, 285, 95 [2] ebisawa, k., maeda, y., kaneda, h., & yamauchi, s. 2001, science, 293, 1633 doi:10.1126/science.1063529 [3] ebisawa, k., tsujimoto, m., paizis, a., et al. 2005, apj, 635, 214 doi:10.1086/497284 [4] ezuka, h. & ishida, m. 1999, apjs, 120, 277 [5] hong, j., schlegel, e. m., & grindlay, j. e. 2004, apj, 614, 508 doi:10.1086/423445 [6] koyama, k., makishima, k., tanaka, y., & tsunemi, h. 1986, pasj, 38, 121 [7] matranga, m., drake, j. j., kashyap, v., & steeghs, d. 2012, apj, 747, 132 doi:10.1088/0004-637x/747/2/132 [8] morihana, k., tsujimoto, m., yoshida, t., & ebisawa, k. 2013, apj, 766, 144m doi:10.1088/0004-637x/766/1/14 [9] nagayama, t., et al., 2003, in society of photooptical instrumentation engineers (spie) conference series, vol. 4841, m. iye & a. f. m. moorwood, 459–464 [10] revnivtsev, m., sazonov, s., gilfanov, m., churazov, e., & sunyaev, a&a, 452, 169r [11] revnivtsev, m., sazonov, s., et al., 2009, nature, 458,1142 doi:10.1038/nature07946 [12] rosati, p., tozzi, p.,et al. 2002, apj, 566, 667 doi:10.1086/338339 [13] schmidt, m. r., & miko lajewska, j. 2003, nearinfrared spectra of a sample of symbiotic stars [14] smith, r. k., brickhouse, n. s., liedahl, d. a., & raymond, j. c. 2001, apjl, 556, l91 doi:10.1086/322992 [15] warwick, r. s., turner, m. j. l., watson, m. g., & willingale, r. 1985, nature, 317, 218 doi:10.1038/317218a0 [16] worrall, d. m., marshall, f. e., boldt, e. a., & swank, j. h. 1982, apj, 255, 111 225 http://dx.doi.org/10.1126/science.1063529 http://dx.doi.org/10.1086/497284 http://dx.doi.org/10.1086/423445 http://dx.doi.org/10.1088/0004-637x/747/2/132 http://dx.doi.org/10.1088/0004-637x/766/1/14 http://dx.doi.org/10.1038/nature07946 http://dx.doi.org/10.1086/338339 http://dx.doi.org/10.1086/322992 http://dx.doi.org/10.1038/317218a0 kumiko morihana, masahiro tsujimoto, ken ebisawa discussion takeshi go turu’s comment: is the fe line a mixture of fe lines (6.4, 6.7 7.0 kev) or 6.7 kev line? the fe line in our grxe spectrum is a mixture of the fe lines (6.4, 6.7, and 7.0 kev). the line center of the fe line is ∼6.7 kev. guainazzi matteo’s comment: radio-quiet agns normally exhibit iron kα lines. this may suggest that your iron free agns could be orimary radio-loud. do you have radio measurements, that could validate this hypothesis? i consider that fe k line of most background agns are red-shifted, because most background agns constituting the grxe are at far distance. 226 introduction analysis and results x-ray observation and source extraction spectral fittings grouping global spectral fittings nir imaging observation and source extraction cross correlation nir spectroscopy discussions & conclusion acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0033 acta polytechnica ctu proceedings 4:33–37, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app impact of the chemical form of in-containment source on fission product release from wwer-1000/v-320 type npp containment during loca adam kecek department of nuclear reactors, faculty of nuclear sciences and physical engineering, czech technical university in prague correspondence: kecekada@fjfi.cvut.cz abstract. nuclear power plant accidents may be followed by a release of fission products into the environment. this release is dependent on several phenomena, such as chemistry, pressure, type of the accident etc. the aim of this paper is to assess the impact of the chemical form of iodine on the fission product release into the environment. keywords: npp, loca, iodine, containment, cocosys, fission products transport, wwer1000/v-320. 1. introduction fission product behaviour and its simulation recently become more important. nuclear power plant accidents in past decades revealed the need of proper definition and calculation of fission product mass entraining the environment. to successfully calculate such phenomena, one has to specify the in-containment source correctly. several different approaches existing up to date vary in the mass, the chemical form or time characteristics. this paper is focused on assessment of the impact of the chemical form of iodine release, which is well known for its extensive range of chemical compounds and reactions within the containment. a fundamental opinion on the complexity of iodine chemistry in the containment during loss-of-coolant accidents can be obtained from fig. 1. used models, codes and brief overview on existing approaches are presented further. 2. description of cocosys the containment code system is a lumped parameter computer code developed and maintained at gesellschaft für reaktorsicherheit (grs mbh). it is used for best-estimate analysis of lightwater reactor containments during severe accidents. the code itself consists of several modules, which calculate all relevant phenomena. the thermal hydraulic (thy) module solves the pressure and temperature build-up history together with mass and energy flow. furthermore it can simulate behaviour of engineering systems like valves, fans, pumps and sprays. a feedback from aerosol and fission products is also considered. the aerosol fission product (afp) module is used to estimate the behaviour and transport of fission products and aerosols within the containment. the core concrete interaction (cci) module provides an investigation of interaction of the molten core with concrete structures during severe accidents. all menfigure 1. scheme of iodine chemistry in npp containment [1]. tioned modules are coupled to provide the desired feedback on relevant phenomena. furthermore, cocosys provides a coupling option with several codes like athlet to simulate primary circuit behaviour or a cfd code for detailed analysis [2]. cocosys has been validated against several validation experiments and is often used in safety analysis of npp concepts and existing power plants including wwer type npps [2]. 3. wwer-1000/v-320 model description the wwer-1000/v-320 type npp containment is a cylindrical shaped building made of pre-stressed 33 http://dx.doi.org/10.14311/ap.2016.4.0033 http://ojs.cvut.cz/ojs/index.php/app adam kecek acta polytechnica ctu proceedings concrete with an inner steel liner. for thermal hydraulic calculations, the inner volume has to be divided into several smaller rooms. the presented model provides an approximation based on 20 volumes, so called nodes, which simulate the real space surrounded by the building. these volumes are connected through junctions providing mass and energy transfer. the walls, floor, ceiling and inner structures are also modeled to provide realistic heat transfer. furthermore the containment spray systems are also modeled. two large nodes (environ and building) simulate the outer volume surrounding the containment. these provide boundary conditions of the problem. the nodalisation of the containment is presented in tab. 1. compartments nodes description e1 environ environment building surrounding buildings iga501 ga501 upper part of the reactor shaft iga301 ga301a lower part of the reactor shaft iga401 ga401a2 spent fuel pools iga701 ga701 reactor hall iasupp cnt-u1 upper section of the building surrounded by the containment wall and steam generator boxes cnt-u2 iasgbox1 sgbox-1 steam generator box 1 iasgbox2 sgbox-2 steam generator box 2 iasmed cnt-m1 middle section of the builcnt-m2 ding at the containment wall iaslow cnt-l1 lower section of the containment building cnt-l2 ga311 ga306-1 ga306-2 ga306-3 iga201 ga201 spray water sump iastl22 tl22p fan system table 1. containment discretization. the simulation of containment leak is maintained by atmospherical junction. this is estimated to follow the limits and conditions of operation of the wwer1000/v-320 type containment. the junction itself is placed in the upper part of the containment connecting the reactor hall and environment. the reason for choosing this position is to simulate shortest and easiest way for ejecting the material from containment into the environment. for calculation of fission product and aerosol behaviour the containment building has to be divided into volumes so-called compartments formed by nodes. thus, compartments provide coarser fragmentation of the containment building. the fragmentation of the containment volume for iodine calculations is presented in tab. 1 and fig. 2. the calculations were done using the afp and thy cocosys computational modules. figure 2. iodine compartment discretization of wwer-1000/v320 npp containment. 4. initial and boindary conditions before the accident, the nuclear power plant is operated at 104 % nominal power. the initiating event is a primary circuit hot leg guillotine rupture at t = 0 s with leak through breaks of equivalent diameter equal to 850 mm (large break loca) into one of the steam generator boxes. moreover, following assumptions are taken into account [3]: • loss of on-site power, • only 1/3 of safety systems in operation (according to conservative approach), • no action expected from operators. the entrainment of fission products into the containment is maintained through the break in the primary hot leg. the ejecting coolant brings the fission products and aerosols released from the core. in accordance with local practice the in-containment source term is based on ref. [4]. following this approach failure of all fuel assembly is expected. in addition, whole fission product mass released into the primary circuit is transferred into the containment. this means that no fission products remain in the primary circuit. the ref. [4] divides the problem of fission product release into two areas, namely normal and core melt. 34 vol. 4/2016 impact of the chemical form of in-containment source the investigated accident does not expect the molten core, so only the pellet-cladding gap inventory is taken into account. the fraction of released fission products is given in tab. 2. fission product pellet-cladding gap fuel melt release release fraction fraction (30–1800 s) (1800–4680 s) i, br 0.05 0.95 xe, kr 0.05 0.35 cs, rb 0.05 0.25 te, sb, se 0.00 0.05 ba, sr 0.00 0.02 ru, rh, pd, mo, 0.00 0.0025tc, co la, zr, nd, eu, 0.00 0.0005nb, pm, pr, sm, y, cm, am ce, pu, np 0.00 0.0002 table 2. fraction of the fission products released during an accident [4]. a research on available in-containment fission product source showed a significant difference between each approach. synchronization efforts on this problem throughout europe pointed out the diversity of the iodine species, the target medium (water, sump or aerosol) and also the fraction of the core inventory released [5]. it has to be mentioned that there is no information on time characteristics of the release. based on this it was decided to test, whether the chemical form and also the target part of the zone has any influence on iodine mass entraining the environment. the time characteristics and fraction released was adopted from ref. [4]. finally, five calculations were done in this sensitivity study, where the distribution of iodine mass varied in chemical form. in addition, the target part of the zone was changed too. overview on all calculations is given in tab. 3. calculation chemical target part iodine mass form of the zone distribution [%] c00 csi atmosphere1 95 i2 4.85 ch3i 0.15 c01 csi atmosphere 0.15 i2 95 ch3i 4.85 c02 csi sump 0.15 i2 95 ch3i 4.85 c03 csi atmosphere 4.85 i2 0.15 ch3i 95 c04 csi sump 4.85 i2 0.15 ch3i 95 table 3. overview on calculated variants. the total mass of fission products in the core is based on values provided in ref. [3]. the total amount of iodine used in this calculation is 22.7 kg. it has to be noted that not all the fission product mass is expected to be radioactive. this study calculates the behavior of the whole mass of iodine entraining the containment. 5. results and comparison the calculations were done for 14 400 s, what equals to 4 hours. this time range was chosen to be representative as the testing calculations showed that the marginal contribution to the mass ejecting the containment during the accident is within first two hours. calculation c04 was neglected due to numerical error. no. description mass [g] deviation from c00 [%] c00 base case (95 % csi, 4.85 % i2, 0.15 % ch3i) a 4.94 × 10−3 0.00 c01 modified iodine mass distribution (95 % i2, 4.85 % ch3i, 0.15 % csi) a 2.33 × 10−2 371.51 c02 modified iodine mass distribution (95 % i2, 4.85 % ch3i, 0.15 % csi) s 4.78 × 10−3 −3.39 c03 modified iodine mass distribution (95 % ch3i, 4.85 % csi, 0.15 % i2) a 1.14 × 10−1 2204.44 c04 modified iodine mass distribution (95 % ch3i, 4.85 % csi, 0.15 % i2) s err err table 4. overview on calculation and results. the trend of the release is in good agreement with all calculations except c03, where almost constant release rate can be observed. the ejected mass in calculations c00 through c02 shows a massive growth of the release within the first hour, then it slows down and in long time view it shows stabilization of the mass entraining the environment. this may be caused by several phenomena, such as sedimentation or wash down of fission products from walls into the sump. the worst case presented in this sensitivity analysis is c03, where the main mass was injected as organic iodine into the atmosphere. it has to be noted that this form is not expected to be a typical contributor to the in-containment source, thus the results seem to be unrealistic. the second worst case is c01, where the iodine in elemental form is injected into the gas part of the zone. this leads to a high release of iodine into the environment. in accordance to ref. [6] the elemental 1the csi aerosol can be injected only into the gas part of the zone. 35 adam kecek acta polytechnica ctu proceedings figure 3. distribution of iodine in the containment (c00–c03, 14 400 s). iodine is one of the main options of iodine mass release into the environment. the base case c00 provided results lower than the two calculations presented before. the mass behaves significantly different as the compound is expected to be in aerosol particle form. the lowest release was calculated for elemental iodine injected into the sump part. this may be caused by high solubility of iodine species. the results are presented in tab. 4 and fig. 4. the distribution within the containment varies for each calculation. in the case of c00 the mass is centralized in the containment sump. this may be induced by the aerosol wash down. the c01 shows significant concentration in the reactor hall. this is probably caused by the chemical form, as the injection was in form of elemental iodine into the atmosphere. furthermore, c02 provides results close to the c00 calculation. the injected iodine might be dissolved in water and then washed down into the containment sump. finally the c03 calculation shows distribution close to the c01. this is again induced by the injection of iodine as a gas. the iodine mass is mainly distributed within the reactor hall. the results of the calculations c00 through c03 at 14 400 s are presented in fig. 3. figure 4. iodine release into the environment (c00–c03, 14 400 s). 36 vol. 4/2016 impact of the chemical form of in-containment source 6. conclusion the presented calculations proved a significant impact of the chemical form of iodine on release into the environment. furthermore the injected phase (water, gas) is also an important factor significantly influencing the ejected mass. future effort should be made on the assessment of the in-containment source, e.g. the time range and the fraction of ruptured fuel assembly. references [1] irsn. the astec software package. http://www.irsn.fr/en/research/ scientific-tools/computer-codes/pages/ the-astec-software-package-2949.aspx. [2] w. klein-heßling, et al. cocosys v2.4 user’s manual. gesellschaft für anlagenund reaktorsicherheit (grs) mbh, 2010. [3] a. kecek. testovací výpočty úniku štěpných produktů z kontejnmentu je temelín v průběhu loca události programem cocosys. tech. rep. z4339, újv, 2016. [4] u.s. nuclear regulatory commission. alternative radiological source terms for evaluating design basis accidents at nuclear power reactors, 2000. regulatory guide 1.183. [5] nuclear fuel behaviour in loss-of-coolant accident (loca) conditions. tech. rep. 6846, nea oecd, 2009. [6] u.s. nuclear regulatory commission. methods and assumptions for evaluating radiological consequences of design basis accidents at light-water nuclear power reactors, 2003. regulatory guide 1.195. 37 http://www.irsn.fr/en/research/scientific-tools/computer-codes/pages/the-astec-software-package-2949.aspx http://www.irsn.fr/en/research/scientific-tools/computer-codes/pages/the-astec-software-package-2949.aspx http://www.irsn.fr/en/research/scientific-tools/computer-codes/pages/the-astec-software-package-2949.aspx acta polytechnica ctu proceedings 4:33–37, 2016 1 introduction 2 description of cocosys 3 wwer-1000/v-320 model description 4 initial and boindary conditions 5 results and comparison 6 conclusion references 103 acta polytechnica ctu proceedings 1(1): 103–107, 2014 103 doi: 10.14311/app.2014.01.0103 the formation of massive stars: from herschel to near-infrared paolo persi1, mauricio tapia2 1institute for space astrophysics and planetology (iaps/inaf) via fosso del cavaliere 100, 00133 roma, italy 2instituto de astronomia,unam, apartado postal 877,ensenada, baja california, cp22830, mexico corresponding author: paolo.persi@iaps.inaf.it abstract we have studied a number of selected high mass star forming regions, including high resolution near-infrared broadand narrow-band imaging, herschel (70, 160, 250, 350 and 500 µm) and spitzer (3.6, 4.5, 5.8 and 8.0 µm) images. the preliminary results of one of this region, iras 19388+2357(mol110) are discussed. in this region a dense core has been detected in the far-infrared, and a young stellar cluster has been found around this core. combining near-ir data with spitzer and herschel photometry we have derived the spectral energy distribution of mol110. finally comparing our h2 and kc narrow-band images, we have found an h2 jet in this region. keywords: star formation circumstellar matter giant molecular clouds infrared. 1 introduction the formation of high-mass stars defined as those with masses greater than 8 msun is still controversial. one of the crucial problem is to understand if high-mass stars can form through(disk) accretion like low-mass stars. such stars reach the zero-age main sequence (zams) still undergoing heavy accretion, and their powerful radiation pressure should halt the infalling material, thus inhibiting growth of the stellar mass beyond about 8 msun (e. g., palla & stahler 1993). recently, various studies have proposed a solution to this problem based on non-spherical accretion and high accretion rates (e. g., mckee & tan 2003; bonnell et al. 2004; kuiper et al. 2010). in addition at difference of low-mass stars, high-mass stars forms in cluster. therefore the comprehension of the massive star formation process requires good observational knowledge of the star-forming environment and of the evolutionary steps through which ob star formation occurs. this can be made combining observations at different wavelengths from near to far-infrared and millimeter wavelengths. thanks to the spitzer and herschel satellites that operate from the mid-ir to the sub-millimeter it is now possible to have a broad observational coverage of these high mass star formation regions. we have selected a number of these regions reported in table 1 with typical characteristics of high-mass star formation(i.e. presence of water and methanol maser sources, radio and millimeter emission , ammonia cores, (molinari et al.1998, molinari et al. 2000). the type high(h) and low(l) reported in table 1 are taken from molinari et al.1996.. we obtained subarcsec resolution nearinfrared broadand narrow-band images of the source of table 1. these observations are compared with far-ir images from the herschel infrared galactic plane survey (hi-gal, molinari et al. 2010) supplemented with spitzer/irac archive images. the observations are described in section 2, while in section 3 we report the preliminary results of iras 19388+2357(mol110). all the results will be discussed in forthcoming papers. table 1: sample of the observed high-mass protostars source type iras α(2000) δ(2000) d l h m s ◦ ′ ” kpc l� mol83 h 18566+0408 18 59 10.0 04 12 15 6.8 1.02 × 105 mol98 l 19092+0841 19 11 37.4 08 46 30 4.5 9.20 × 105 g45.07+0.13 19110+1045 19 13 22.6 10 50 53 9.7 1.42 × 105 g45.47+0.13 19 14 08.3 11 12 32 6.0 3.80 × 105 mol110 h 19388+2357 19 40 59.4 24 04 39 4.3 1.48 × 104 103 http://dx.doi.org/10.14311/app.2014.01.0103 paolo persi, mauricio tapia 1.1 near-infrared images near-infrared images through narrow-band h2 (λo = 2.122 µm, ∆λ = 0.032 µm ) and kcont ( λo = 2.270 µm, ∆λ = 0.034 µm) filters, as well as through standard broad-band jhks filters, were collected on the nights of 2008 july 12 and 14 using the near infrared camera spectrometer (nics) attached to the 3.58m telescopio nazionale galileo (tng) at the observatorio del roque de los muchachos on la palma island. nics has a hgcdte hawaii 1024 ×1024 array and was used in the sf (small field) configuration with a plate scale of 0.13 arcsec/pixel. in each band, 9 dithered frames spaced by 10 arcsec were taken and coadded, for total on-source integration times of 630 s, 540 s and 360 s for j, h, and ks, respectively. the total integration time for each of the narrow-band (h2 and kcont) filters was 1170 s. all images were calibrated using photometric standard stars from hunt et al. (1998) and persson et al. (1998). the measured fwhm of the point-spread function (psf) is between 0.6 arcsec and 0.8 arcsec. jhk photometry was obtained using daophot (stetson 1987) within iraf in the standard way, with an aperure of 1 arcsec. for the crowded regions, we used the psf procedure, also within iraf. 1.2 hi-gal images hi-gal is a herschel open time key-project (molinari et al.2010) aiming at mapping the galactic plane with the pacs (70 and 160 µm, poglitsch et al. 2010) and spire (250, 350, and 500 µm, griffin et al. 2010) photometers on board the herschel satellite (pilbratt et al. 2010). our target were observed by spire+pacs in parallel mode at a scan speed of 60 arcsec/s. the data were reduced using the hi-gal standard pipeline (traficante et al. 2011). the images have pixel sizes 3.2 arcsec, 4.5 arcsec, 6 arcsec, 8 arcsec, and 11.5 arcsec , at 70, 160, 250, 350, 500 µm respectively. from the images, we performed the source extraction and photometry using the curvature threshold extractor package (cutex, molinari et al. 2011). 1.3 spitzer/irac archive images flux-calibrated images of our regions were retrieved from the glimpse (benjamin et al. 2003, churchwell et al. 2009) survey taken at 3.6, 4.5, 5.8 and 8 µm with irac on board the spitzer space observatory . the flux densities of the mid-infrared counterparts of our sources were extracted from the glimpse catalogue. 2 iras19388+2357(mol110) iras 19388+2357 is associated with h2o maser emission (palla et al. 1991; brand et al. 1994), a radio source (hughes & macleod 1994; molinari et al. 1998) and dense molecular gas traced in nh3 by molinari et al. (1996) who renamed as mol110. methanol maser was also detected from this source (schutte et al. 1993; slysh et al. 1994). zhang et al. (2005) detected a co outflow. the centroid of their co emission is approximately 29 arcsec south of the iras position. finally beltran et al.(2006) detected a dense core at 1.2mm with a mass of 167 msun. the presence of a uchii region, of water and methanol maser, confirm that mol110 is an high-mass star forming region. figure 1 shows our ks-band image including the positions of the mentioned sources. figure 1: ks-band image of iras19388+23657. the plus indicates the position of the iras source, while the open circle and the cross give the positions of the 6cm radio continuum and the msx source respectively. the contours represent the 1.2mm emission. a point-like far-infrared sources has been detected in the hi-gal herschel images from 70 to 500 µm at the position α2000 = 19 h40m59.2s, δ2000 = +24 ◦ 04’ 44.9”. we have analyzed our jhks and spitzer images within an area of approximately 20” × 20” around this position. figure 2 reports the color-coded jhks and spitzer images of this area. within the herschel beam we found several very red sources. we have obtained the jhks photometry of these sources around the herschel peak position. from this photometry, we have obtained the j − h versus h − ks diagram illustrated in figure 3. more than 14 objects show significant near-infrared excess, suggesting the presence of a young stellar cluster in this region. the positions of these sources are marked in fig.2 (left panel). at least six of these sources are identified with the mid-ir spitzer sources (see fig.2 right panel). 104 the formation of massive stars: from herschel to near-infrared figure 2: (left panel) jhks color-coded image of iras19388+2375 obtained combining the j (blue), h (green), and ks (red) individual images the contours show the 70 µm herschel observation. the symbol (+) marks the positions of the sources with near-ir excess, while the crosses indicate sources not detectect in j but with h-ks greater than 3(right panel) color-coded spitzer image obtained combining the [3.6] (blue), [4.5] (green), and [8.0] µm (red) individual irac images. the central position of the two images is α2000 = 19 h40m59.1s, δ2000 = +24 ◦ 04’ 45.4”. north is at the top and east to the left. figure 3: j-h versus h-ks diagram relative to a region of 20arcsec around the iras position. on the base of the position coincidence, we have identified the far-ir source with a near-ir and spitzer source. at the same position a source has been detected also with the wise satellite. combining data from near-ir to millimeter spectral region we have constructed the spectral energy distribution (sed) of mol110 reported in figure 4. the sed has been fitted with the infalling envelope+disc+central source radiation transfer model described by robitaille et al. (2006) by using the fitting tool of robitaille et al. (2007). the parameters of the model that best fit the sed are reported in table 2. figure 4: spectral energy distribution (sed) of mol 110 the observed luminosity of 1.13 104 l� correspond to that of a b1-2 zams star reddened by 50.2 magnitudes of extinction in v. 105 paolo persi, mauricio tapia figure 5: (left panel) h2 narrow-band image of iras 19388+2357. (right panel) narrow-band kc image of the same region. north is at the top and east to the left. table 2: physical parameters of mol 110 derived from the robitaille et al. (2007) model. stella mass (msun) 12.94 stellar temperature (k) 12262 envelope accretion rate (msun/yr) 3.61 10 −3 envelope outer radius (au) 1.0 105 envelope cavity angle (deg) 25.1 disk mass (msun) 1.4910 −1 disk outer radius (au) 23.1 disk accretion rate (msun/yr) 5.18 10 −6 av 50.2 d(kpc) 4.7 lbol 1.13 10 4 l� from the comparison of our narrow-band images centered on the h2 (λo = 2.122 µm), and nearby continuum kcont ( λo = 2.270 µm), we have found an h2 jet at the position α2000 = 19 h40m59.7s, δ2000 = +24 ◦ 04’ 49.0” . a nearby source at the position α2000 = 19h40m59.5s, δ2000 = +24 ◦ 04’ 47.6” with near-ir excess could be the young stellar object (yso) driving the observed outflow. this is illustrated in figure 5. similar observations obtained by varricatt et al. (2010) report three different h2 jets in the region not detected in our images. the positions of these h2 jets are very far from the co outflow observed by zhang et al. (2005), indicating that another yso is responsible of driving this outflow. 3 conclusions in order to understand the physical processes that involve high massive star forming regions, the comparison of observations at different wavelengths from near-ir to millimeter are fundamental. we have here reported an example of this combined analysis including near-ir images, spitzer data from 3.6 to 8 µm and herschel images in five bands from 70 to 500 µm, relative to the star forming region iras 19388+2357. from this analysis the following conclusions can be made: 1) a very dense and cold core has been detected from the farir herschel images, in proximity of the iras source. 2) within the dense core the near-ir images show the presence of a young stellar cluster of at least 15 members in a radius of 20 arcsec. 3) the far-ir peak has been identified with a bright spitzer and near-ir source. combining the photometry from 1.25 µm to 1.2mm, we have derived its spectral energy distribution(sed). the measured total luminosity indicates that the source is a 106 the formation of massive stars: from herschel to near-infrared b1-2 zams with av =50.2 4) finally, the narrow-band image centered on the h2 line at 2.122 µm shows the presence of an h2 jet in proximity of the iras source. references [1] beltran, m. t., brand, j., cesaroni, r., et al.: 2006, a&a, 447, 221 [2] benjamin, r. a., churchwell, e., babler, b. l., et al.: 2003, pasp, 115, 953 doi:10.1086/376696 [3] bonnell, i. a., vine, s. g., bate, m. r.: 2004, mnras, 349, 735 doi:10.1111/j.1365-2966.2004.07543.x [4] brand j., et al.: 1994, a&as, 103, 541 [5] churchwell, e., babler, b. l., meade, m. r., et al.: 2009, pasp, 121, 213 doi:10.1086/597811 [6] griffin, m. j., abergel, a., abreu, a., et al.: 2010, a&a, 518, l3 [7] hughes v. a., macleod g. c.: 1994, apj, 427, 857 [8] hunt, l. k., mannucci, f., testi, l., et al.: 1998, aj, 115, 2594 [9] kuiper, r., klahr, h., beuther, h., henning, t.: 2010, apj, 722, 1556 doi:10.1088/0004-637x/722/2/1556 [10] mckee, c. f., tan, j. c.: 2003, apj, 585, 850 doi:10.1086/346149 [11] molinari s., brand j., cesaroni r., palla f.: 1996, a&a, 308, 573 [12] molinari, s., brand, j., cesaroni, r., palla, f., palumbo, g. g. c.: 1998, a&a,336, 339 [13] molinari, s., brand, j., cesaroni, r., palla, f.: 2000, a&a, 355, 617 [14] molinari, s., swuyard, b., bally, j., et al.: 2010, a&a, 518, l100 [15] molinari, s., faustini, f., schisano, e., et al.: 2011, a&a, 530, a133 [16] palla f., brand j., cesaroni r., comoretto g., felli m.: 1991,a&a, 246, 249 [17] palla, f., stahler, s. w.: 1993, apj, 418, 414 [18] persson, s. e., murphy, d. c., krzeminski, w., et al.: 1998, aj, 116, 2475 [19] pilbratt, g. l., riedinger, j. r., passvogel, t., et al.: 2010, a&a, 518, l1 [20] poglitsch, a., waelkens, c., geis, n., et al.: 2010, a&a, 518, l2 [21] robitaille, t. p., whitney, b. a., indebetouw, r., et al.: 2006, apjs, 167, 256 [22] robitaille, t. p., whitney, b. a., indebetouw, r., wood, k.: 2007, apjs, 169, 328 [23] schutte a. j., van der walt d. j., gaylard m. j., macleord g. c.: 1993, mnras, 261, 783 [24] slysh v. i., dzura a. m., valtts i. e., gerard e.: 1994, a&as,106, 87 [25] stetson, p. b.: 1987, pasp, 99, 191 [26] traficante, a., calzoletti, l., veneziani, m., et al.: 2011, mnras, 416, 2932 [27] varricatt, w. p., davis, c. j., ramsay, s., todd, s. p.: 2010, mnras, 404, 661 bibitem zhang q., hunter t. r., brand j., et al.: 2005, apj, 625, 864 107 http://dx.doi.org/10.1086/376696 http://dx.doi.org/10.1111/j.1365-2966.2004.07543.x http://dx.doi.org/10.1086/597811 http://dx.doi.org/10.1088/0004-637x/722/2/1556 http://dx.doi.org/10.1086/346149 introduction near-infrared images hi-gal images spitzer/irac archive images iras19388+2357(mol110) conclusions 174 acta polytechnica ctu proceedings 2(1): 174–177, 2015 174 doi: 10.14311/app.2015.02.0174 application of total variation minimization to doppler tomography m. uemura1, t. kato2, d. nogami3, r. mennickent4 1hiroshima astrophysical science center, hiroshima university, kagamiyama 1-3-1, higashi-hiroshima, hiroshima 739-8526, japan 2department of astronomy, kyoto university, kitashirakawa-oiwake-cho, sakyo-ku, kyoto 606-8502, japan 3kwasan and hida observatories, kyoto university, yamashina-ku, kyoto 607-8471, japan 4universidad de concepción, departamento de astronomı́a, casilla 160-c, concepción, chile corresponding author: uemuram@hiroshima-u.ac.jp abstract we have developed a new model of the doppler tomography using total variation minimization (dttvm). we demonstrated that this method can reconstruct localized and non-axisymmetric profiles possibly having sharp edges in the doppler map. we apply this model to the real data of the dwarf nova, wz sge in superoutburst. dttvm can reproduce the observed spectra with a high precision, while the previous models fail to reproduce localized sources. keywords: cataclysmic variables dwarf novae optical spectroscopy individual: wz sge. 1 introduction doppler tomography (dt) is a widely-used method in the field of cataclysmic variables (cvs) [1]. dt is an ill-posed inverse problem in most cases. a filtered backprojection method provides a way to solve the problem. it masks high frequency components in the doppler maps in order to uniquely determine the solution. the other way is to introduce regularization terms in addition to the likelihood term. the maximum entropy method (mem) has been used to regularize the dt problem. the dt with mem determines the solution by maximizing the information entropy over the default image. the comparison between the back-projection method and mem is reported in [2]. those dt methods tend to smeared out the localized and sharp-edge profiles in the doppler map. on the other hand, such profiles are expected to appear in accretion disks of cvs, such as spiral shocks and hot spot. it has been reported that there are significant residuals between the observed and model spectra that are obtained from dt (e.g. [3]). it is, however, unclear if the residuals indicate real emission components that are not reproduced by the model. in this paper, we introduce a new dt by total variation minimization (tvm), which is described in [4]. tvm has received attention in the field of image reconstruction because it can reconstruct sharp-edge features by making image sparse in its gradient domain. 2 doppler tomography by total variation minimization 2.1 model suppose there are m data given by d = {d0,d1, · · · ,dm}. the i-th element of d, di, is the flux at a radial velocity vr,i and phase φi. also suppose n × n = n points in the doppler map given by s = {s0,0,s0,1, · · · ,sn,n}. the j-th element in the map has the coordinate of (vx,j,vy,j) and intensity of sj. the goal of dt is to estimate ŝ from the observation d according to ŝ = arg min s { 1 m ‖d−as‖22 + λ n φ(x) } . (1) the first term in the right side is a least-square term, and the second term is a regularization term which consists of a hyperparameter λ and a function φ. a is a m × n matrix whose element depends on the phase, radial velocity, and instrumental response. for the response function, we assumed a gaussian profile with a variance σ2, where σ corresponds to the velocity resolution of the observed spectra. an element of a, aij is then given by aij = 1 √ 2πσ2 exp { − (vr,i −vr,ij)2 2σ2 } . (2) tvm provides a method to solve ill-posed problems. 174 http://dx.doi.org/10.14311/app.2015.02.0174 application of total variation minimization to doppler tomography the isotropic total variation φtvm,iso is defined as φtvm,iso(s) = ∑ i,j √ (si+1,j −si,j)2 + (si,j+1 −si,j)2. (3) φtvm,iso is used as φ in equation (1) in our model. we use the twist algorithm to estimate ŝ in equation (1) [5, 6].1 the calculation code of dttvm can be available at our web site.2 2.2 demonstrations using artificial data we show an example of the results of dttvm using artificial data sets in figure 1. first, we made the assumed doppler maps, then simulated observations of line-profile variations, and reconstructed the maps using dttvm. we calculated three doppler maps from the simulated spectra with different values of λ in equation (1). we assumed the original doppler map containing three spot structures, as shown in the upper leftmost panel of figure 1. the upper two spots have sharp-edge and flat-top profiles of different size and intensity. the lower left spot has a gaussian profile. dttvm reproduced the original map in all three cases. in case (a), the edges of the flat-top spots are slightly smeared out compared to cases (b) and (c). in addition, the peak intensity of the gaussian spot is underestimated in case (a). these smearing effects are less prominent for smaller values of λ. the model spectra and residuals are also shown in figure 1. the residuals are smaller for smaller λ. these results demonstrate that dttvm can reconstruct both sharp-edge and smooth profiles in a doppler map. 2.3 application to the data of wz sge we apply dttvm to the data of the dwarf nova, wz sge taken on the 10th day of the superoutburst in 2001 [7]. figure 2 presents the doppler maps and model spectra. the results were computed using log10 λ of (a) −2.2, (b) −3.2, and (c) −5.2. all doppler maps exhibit a disk feature having non-axisymmetric distribution of intensity. the peak intensity of the localized bright regions is lower in case (a) than the others. as a result, the residuals are large in case (a), compared with the other cases. the model spectra reproduce the observed one without any prominent residuals in cases (b) and (c). figure ?? shows the result of the doppler tomography using the same data of wz sge calculated with mem. we used the calculation code presented in [8].3 the calculation code defines α as a hyperparameter of the model. cases (a), (b), and (c) in figure ?? were calculated with log α = −1.0, −2.0, and −3.0, respectively. in cases (a) and (b), smooth two-armed structures appear on a weak sign of a disk feature. in case (c), patchy structures appear, but they are evidently artifacts. in all cases of (a), (b), and (c), there are significant residuals of the spectral data at −500 km s−1 for phase 1.75 and at +500 km s−1 for phase 1.60. both features have counterparts at phases shifted by 0.5, which indicate they are rotating components with the orbital period. these rotating residuals definitely arise from localized structures in the doppler map, which can be successfully reconstructed by dttvm: the former one corresponds to the feature at (vx,vy) ∼ (0, 500) and the latter one to the lower right region in figure 2. 3 summary as demonstrated above, dttvm can reconstruct both sharp-edge and smooth profiles in a doppler map. hence, this model is advantageous for the reconstruction of highly localized or edge structures in the doppler map, such as the emission from the secondary star, hot spot, or shock wavefront. the doppler map of wz sge obtained by dttvm is surprisingly different from that built with the mem method. the difference in the regularization term could cause the different maps especially when the data sample is small. dttvm can reproduce the observed spectra with a high accuracy, even in the case that the mem method generate significant residuals in spectra. this indicates that tvm is probably a suitable regularization method for real doppler map, compared with mem. acknowledgement this work was supported by jsps kakenhi grant number 22540252 and 25120007. we appreciate comments and suggestions from dr. shiro ikeda. rem acknowledges support by fondecyt grant 1110347 and the basal centro de astrofisica y tecnologias afines (cata) pfb–06/2007. 1〈http://www.lx.it.pt/˜bioucas/twist/twist.htm〉 2〈http://home.hiroshima-u.ac.jp/uemuram/dttvm/〉 3〈http://www.mpa-garching.mpg.de/˜henk/〉 175 m. uemura et al. references [1] marsh, t. r. & horne, k. 1988, mnras, 235, 269 doi:10.1093/mnras/235.1.269 [2] marsh, t. r. 2001, in astrotomography, indirect imaging methods in observational astronomy, ed. h. m. j. boffin, d. steeghs, & j. cuypers (springer-verlag), 1 [3] tappert, c., mennickent, r. e., arenas, j., matsumoto, k., & hanuschik, r. w. 2003, a&a, 408, 651 [4] uemura, m., kato, t., nogami, d., & mennickent, r. e. 2013, pasj, submitted [5] bioucas-dias, j. m. & figueiredo, m. a. 2007b, in image processing, 2007. icip 2007. ieee international conference on vol. 1 (ieee), pp i–105 [6] bioucas-dias, j. m. & figueiredo, m. a. 2007a, image processing, ieee transactions on, 16, 2992 [7] nogami, d. & iijima, t. 2004, pasj, 56, 163 [8] spruit, h. c. 1998, arxiv:astro-ph/9806141 discussion dmitry kononov: 1. it’s nice that your method can restore flat features. 2. it’s in a sense incorrect to compare two results of inverse-problem solving methods if you have no physical basics under the structures you are trying to explain. makoto uemura: as you pointed out, both tvm and mem have no physical meaning. but, i think the comparison still makes a sense. we can select a better model based on residuals between observations and model. in the case of wz sge, the residuals obtained with our method are much smaller than those with mem. the mem method fails to reproduce the localized structures of the line profile, while our method can reproduce them. dmitry bisikalo: is it possible to use tvm in 3d doppler tomography? makoto uemura: yes, it is. the 3d doppler tomography is also a linear problem, so it is easy to apply the method to it. vx (km s−1) v y (k m s −1 ) −1000 0 +1000 − 1 0 0 0 0 + 1 0 0 0 original − 1 5 1 0 −1000 0 +1000 − 1 0 0 0 0 + 1 0 0 0 (a)logλ=−1.0 −1000 0 +1000 (b) logλ=−2.0 −1000 0 +1000 (c) logλ=−3.0 rv (km s−1) o rb ita l p h a se −2000 0 +2000 0 .0 0 .2 0 .4 0 .6 0 .8 1 .0 −2000 0 +2000 0 .0 0 .2 0 .4 0 .6 0 .8 1 .0 0 +2000 0 +2000 0 +2000 0 +2000 0 +2000 figure 1: results of dttvm using an artificial map having three spots. the artificial map is depicted in the upper left-most panel. the simulated spectra from the map is shown in the lower left-most panel. three sets of results are shown with different λ, which is indicated by the top of the figure. for each set, the estimated doppler map is shown in the upper panel. the lower left and right panels are the model spectra and residuals between the original and model spectra, respectively. 176 http://dx.doi.org/10.1093/mnras/235.1.269 application of total variation minimization to doppler tomography vx (km s−1) v y (k m s −1 ) −1000 0 +1000 − 1 0 0 0 0 + 1 0 0 0 (a) logλ=−2.2 vx (km s−1) v y (k m s −1 ) −1000 0 +1000 − 1 0 0 0 0 + 1 0 0 0 (b) logλ=−3.2 vx (km s−1) v y (k m s −1 ) −1000 0 +1000 − 1 0 0 0 0 + 1 0 0 0 (c) logλ=−5.2 rv (km s−1) o rb ita l p h a se −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 figure 2: results of dttvm for the data of wz sge. upper panels: the doppler maps. lower left and right panels: the model spectra and residuals between the model and observed spectra, respectively. the three sets of the panels were calculated with different λ, which are indicated in the figure. vx (km s−1) v y (k m s −1 ) −1000 0 +1000 − 1 0 0 0 0 + 1 0 0 0 (a) logα=−1.0 vx (km s−1) v y (k m s −1 ) −1000 0 +1000 − 1 0 0 0 0 + 1 0 0 0 (b) logα=−2.0 vx (km s−1) v y (k m s −1 ) −1000 0 +1000 − 1 0 0 0 0 + 1 0 0 0 (c) logα=−3.0 rv (km s−1) o rb ita l p h a se −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 rv (km s−1) −2000 0 +2000 1 .0 1 .5 2 .0 2 .5 figure 3: as in figure 2, but for the mem case. 177 introduction doppler tomography by total variation minimization model demonstrations using artificial data application to the data of wz sge summary acta polytechnica ctu proceedings doi:10.14311/app.2016.3.0065 acta polytechnica ctu proceedings 3:65–70, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app load bearing capacity of the glass railing element elizabeta šamec∗, domagoj damjanović, joško krolo university of zagreb, faculty of civil engineering, fra a. k. miošića 26, zagreb, croatia ∗ corresponding author: elizabeta.samec@gmail.com abstract. in this paper some basic physical and mechanical properties of glass as structural material are presented. this research is about specifically manufactured glass railing element that will be a part of a pedestrian bridge construction in zagreb, croatia. load bearing capacity test of the glass railing element is conducted within the faculty of civil engineering in zagreb and obtained experimental results are discussed and compared to the ones provided by the numerical model. taking into account the behaviour of laminated glass and results of experimental and numerical testing, glass railing element can be regarded as safe. keywords: glass railing element, load bearing capacity test, wind load, modulus of elasticity, numerical model. 1. introduction when we have glass elements that are specifically manufactured for construction projects it is highly important to perform laboratory testing in order to confirm their load capacity. necessity for laboratory testing follows from potential unexpected breakdown under the load due to special properties of the glass as a construction material. unlike steel, plastic and wood, glass has no plastic properties and consequently no yield point [1]. because of the property of pure elasticity with brittleness the glass cannot be permanently deformed by load and it fails without warning as shown on a stress-strain curve (fig. 1). failure always happens in tension because it has high compressive strength [2]. the cause of the glass failure is not necessarily concentrated impact; fracture may occur due to bending stress, thermal stress or deformation. a fracture which occurs will depend on the number of surface defects, the duration of the load, the size of the tensed surface and the stress level. by increasing the dimensions of the glass elements the number of imperfections will increase as well [3]. this kind of behaviour is undesirable for a construction material. despite possible problems with brittle behaviour and unexpected breakdown, application of structural glass has undergone remarkable development in the last twenty years. the opportunities and attractiveness that this type of material provides have been recognized and glass has become more than “window material”. the subject of this research is the glass railing element designed for the pedestrian bridge construction planned to be built in zagreb, croatia. usually for tempered glass guardrails, live loads caused by use and wind loads need to be considered. other loads such as a permanent load, snow load, and seismic load are insignificant because of their low range compared to the overcharge caused by use or wind [4]. in this case, the wind is also the main load on the glass element, figure 1. stress-strain diagram for steel, glass and plastic [2]. so it is necessary to test the bearing capacity of the glass railing element under the influence of the wind. in most cases glass will behave entirely elastic until the moment of fracture and this kind of behaviour is expected to be confirmed by experimental analysis. the literature review in [1] shows that earlier research has been conducted on behaviour, strength, analytical and numerical modelling of laminated glass with pvb but still the knowledge required to use laminated glass as a load bearing element in every-day construction, is not sufficient. the aim of this paper is to gain a deeper knowledge of the behaviour of laminated glass by experimental investigations and by numerical model simulation. to pursue the proposed study, load bearing capacity test was performed on the glass railing element in the scale 1:1. corresponding finite element numerical models were created in the software sap 2000 to discuss and compare displacements in the points most affected by the wind influence. the main question that needs to be answered is whether the glass railing element can be regarded as safe to install on the pedestrian bridge. therefore, safety factor is calculated and expected 65 http://dx.doi.org/10.14311/app.2016.3.0065 http://ojs.cvut.cz/ojs/index.php/app e. šamec, d. damjanović, j. krolo acta polytechnica ctu proceedings figure 2. bridge plan. figure 3. glass railing element with positions of attachment details. behaviour under the wind load is tested on a life-size sample of the glass railing element. 2. bridge project details 2.1. bridge construction the future pedestrian bridge is planned to be built in miramarska street in zagreb. the railing elements are going to be put along the bridge to protect the pedestrians from the influence of wind and to ensure their safety. the plan of the future bridge is shown in fig 2. 2.2. railing element the glass element consists of two 10 mm thick tempered glass panes bonded by interlayer (polyvinyl butyral-pvb). the stiffness of laminated glass depends on the shear strength of the panes and the pvb foil. factors that influence the behaviour of these connections are load duration, intermediate layer thickness, temperature and position of the intermediate layer with respect to the center of gravity. the behaviour of laminated glass is different under the long-term and short-term load. under the influence of short-term load, laminated glass acts as a composite carrier. under the influence of long-term load, the load is distributed on two panes due to their rigidity because of the deformation of the interlayer. in the case of a breakdown, glass fragments remain linked to the pvb layer. therefore, laminated tampered glass figure 4. detail 1: spider fitting (up), detail 2: inox clamps connecting the element to the main steel profile (down). is chosen for the railing element of the pedestrian bridge. the dimensions of the railing element are 213.4 x 107.5 cm. the element is attached to the bridge construction at four points. as you can see in the fig. 3 and fig. 4, the element is put in the inox glass clamps on the bottom, and on the handlebar section it is fixed to the handlebar with two spider fittings. 2.3. wind load as mentioned in the introduction, the wind load is authoritative load and calculation of the design value used in laboratory and numerical analysis is further described below. the design value is calculated according to croatian valid norm for wind load hrn en 1991-1-4:2012/ na2012. depending on the sensitivity of construction on the dynamic impulse, two procedures exist to calculate wind load: • simplified procedure: applies to structures that are insensitive or moderately sensitive to dynamic impulse, • detailed procedure: applies to structures sensitive to dynamic impulse which have the value of the dynamic coefficient greater than 1.2. 66 vol. 3/2016 load bearing capacity of the glass railing element figure 5. load bearing capacity test. figure 6. positions of measurement points (up) and load distribution (down). for the construction of the pedestrian bridge in zagreb the simplified method is applied since the construction is not sensitive to dynamic impulse. simplified calculation implies that the wind action is taken as a replacing static load. for the bridges, the wind pressure is calculated from the forces in all horizontal directions. to determine the wind load on the glass railing element according to hrn en 1991-1-4:2012/ na2012 the following parameters are required: • basic wind speed: vb = 20 m/s • height above the ground: h = 5 m • field category: iv. the design value of the wind load determined by the project designer is 1 kn/m2. 3. load bearing capacity test the load bearing capacity of the element (in the scale 1:1) was tested on one sample until breakdown in the structural testing laboratory operating at the figure 7. breakdown of the glass railing element. faculty of civil engineering in zagreb. the testing was conducted on the universal testing machine zwick z600. during the test, displacements and strains were measured using lvdt sensors (p) and strain gauges (t) as shown in fig. 6 (up). the sample was tested in the horizontal position and the wind load was simulated with two types of loading. the first type is a continuous load (1 kn/m2) on the surface between the supports (107.5 × 120.3 cm) which was applied with 12 weights during the whole test. the second one is a linear load (qpok) applied on the console part with the help of the previously mentioned machine zwick z600. the two types of loading can be seen on the fig. 5. the bending moment at the handlebar section provoked by the design value of wind load (f=1 kn/m2) on the surface of the console part of the glass element (fig. 6) is calculated as follows: m = f ∗ a22 2 = 1 ∗ 0.9312 2 = 0.433 kn m/m. (1) the linear load (qpok) applied in the experiment must cause bending moment at the support (handlebar section) equivalent to the one caused by the assumed wind force calculated above: m = qpok ∗ a2/2, (2) qpok = m a2/2 = 0.433 0.466 = 0.93 kn/m. (3) knowing the required value of liner load (qpok) concentrated force (fpok) on the steel element is according to fig. 6 (down): fpok = qpok ∗ b = 0.93 ∗ 1.075 = 1.0 kn. (4) in this way continuous load on the whole surface of the railing element that would be provoked by wind 67 e. šamec, d. damjanović, j. krolo acta polytechnica ctu proceedings figure 8. bending test for glass canopy specimens [5]. sample strain [‰] e [mpa] 1 0.60 38 833 2 0.58 40 167 3 0.56 41 607 table 1. results of the bending test at the glass canopy specimens [11]. is simulated. by using a machine to apply fpok we were able to continually increase the force until the breakdown of the sample. the linear load is applied by increments of 0.5 kn until the force value of 2.0 kn, after which it is applied continuously until breakdown. after each load stage, the element is unloaded to determine the existence of any permanent displacements and strains. when the force reached the value of 6.55 kn, breakdown started rapidly expanding through the sample (fig. 7). the safety factor of the glass element can be calculated as the ration between failure load and design load. sf = 6.55 kn 1.00 kn = 6.55 . (5) 4. simulation of load bearing capacity test in sap 2000 as well as the laboratory test, a numerical model was made in sap 2000. a numerical modelling of laminated glass by finite element method is complex, mainly because of the very thin interlayer in comparison with other dimension and because of the large difference in modulus of elasticity of glass ply and interlayer material, especially for pvb [1]. the failure of a laminated glass sheet can be subdivided in five phases [5]: (1.) elastic behavior of the glass plies. (2.) the first glass ply is broken, the other is still intact; the interlayer is not damaged. (3.) the second glass ply fails; the interlayer behaves elastically. (4.) the interlayer behaves plastically; the splinters are kept together by the interlayer. (5.) the interlayer fails by reaching its failure strength or by cutting from the splinters. while phase (1.) can be modelled with analytical or numerical methods, phases (2.) to (5.) are more figure 9. numerical model in sap2000 for fpok = 1.00 kn (up) and fpok = 6.55 kn (down). (u-displacement, r-rotation, 1-x axis, 2-y axis, 3-z axis) complex to simulate. as shown in [5] several material models based on finite elements can be found in the literature to simulate the failure of the glass as well as of the interlayer. models with one shell element through the thickness use layered materials with integration points over the thickness. in [6] a material model for the glass which allows a two dimensional failure is presented. a combination of two shells and one solid element through the thickness can be found in [7]. authors [8], [9] and [10] present 3d models with solid elements which allow using a detailed material law for the interlayer. the aim of the paper is just to compare displacement results in points most affected by wind, so it is concluded that a single thin plate model describes the behaviour of laminated glass for the current application well enough, and it is not expensive in terms of time and problem size. since the test to determine mechanical properties was not performed for this specific railing element, modulus of elasticity was chosen based on the previously known test results. these results are obtained in the same laboratory during the test conducted for the construction of glass canopy [11]. for the glass canopy analysis a bending test was performed on three laminated tempered glass specimens consisting of two 10 mm thick tempered glass panes as shown in fig 8. during the test, strains were measured using strain gauge in the middle of the sample. from the obtained data, modulus of elasticity was calculated using hooke's law. results of this testing are summarized in the table 1, and since the railing element also consists of two 10 mm thick tempered glass panes, the chosen value of modulus of elasticity for this numerical model is e= 40 000 mpa. supports are defined in reliance to fig. 4 and they have prevented displacements in all three directions (x, y and z) while the rotation angle is free around all three axes [12]. the first numerical analysis was made for the load of 1 kn/m2 on the surface between the supports and linear load qpok corresponding fpok= 1.00 kn as it is 68 vol. 3/2016 load bearing capacity of the glass railing element loading experiment sap 2000 force displacement u3 fpok = 1.00 kn 8.00 mm 6.70 mm fpok = 6.55 kn 58.54 mm 60.80 mm table 2. comparison of experimental and numerical displacement results. shown in the fig. 6 (down) and described earlier. the second numerical analysis was conducted in the same way for the maximum load at breakdown fpok= 6.55 kn obtained during the experimental testing. models and given results are shown in fig. 9. 5. results results obtained by the load bearing capacity test are processed and shown in force-strain (fig. 10) and force-displacement diagram (fig. 11). from the force-strain diagram obtained during the experimental analysis, we can confirm that the glass behaved linearly elastic without any damages until breakdown. after all phases of loading, the residual strains were within the boundaries of 4% and the residual displacements within the limits of 10%. even though the displacements were measured during the test at 6 points (p1-p6), the most interesting points to compare obtained experimental and numerical results are points p1 and p2. in those points wind will cause the biggest displacement since they are the top points of the free console part (fig. 6 (up)). in the numerical model there will be no difference between the displacement of the point p1 and p2 so the point p2 is chosen due to an easier display of numerical results. the comparison of the displacement results obtained by numerical and experimental analysis at the measurement point p2 is given in table 2. in fig. 8 displacement gained from numerical model is marked with u3. from the table we can see that the obtained experimental and numerical results differ by less than 15%. as already discussed in the previous section, the simple single plate model is chosen. the thin plate model does not take into account a fact that laminated glass is three layer composite, instead the modulus of elasticity is defined for whole element. the test to determine the mechanical parameters was not performed, so the chosen value of modulus of elasticity based on the comparison of previously known test results can differ for two main reasons. the first reason is that even though the thickness of the glass panes was the same in both cases, different glass is used in manufacturing canopy and railing elements. secondly, we do not know exactly what impact interlayer has on the modulus of elasticity of composed element. figure 10. force-strain diagram. figure 11. force-displacement diagram. 6. conclusion the combination of hard glass with a soft interlayer limits the use of laminated glass as structural load carrying element due to difficult behaviour prediction. glass has a modulus of elasticity several thousand times larger than the interlayer material, which makes the behaviour and the modelling complex. with the simple single thin plate model used in this study, where modulus of elasticity for the whole laminated glass sheet was chosen based on the previous similar testing, gained results for displacement differ from experimental ones by less then 15%. the results show that for studies which include stress-failure models, taking into account the interaction behaviour of glass with interlayer is compulsory. it is also important to perform the test to determine mechanical parameters for each composed glass element because their accuracy can considerably influence the quality of the numerical model. in future studies it is necessary to observe the influence that thickness of interlayer has on the modulus of elasticity of laminated glass elements. the proper understanding of the interaction behaviour will enable wider usage of laminated glass as load carrying element. the safety factor of tested railing element reached in regard to the design wind load is k = 6.55. the glass railing element can endure six times higher load than anticipated so we can conclude that there are high 69 e. šamec, d. damjanović, j. krolo acta polytechnica ctu proceedings reserves in the bearing capacity of tested glass element. since the test was performed at the room temperature and pvb is highly temperature dependent material, by taking into account the weather conditions on the bridge we can expect the safety factor to be lower. considering the behaviour of laminated glass during the test and results of experimental testing compared to numerical ones, glass railing element can be regarded as safe. list of symbols vb basic wind speed [m/s] fpok concentrated force in the experiment [kn] f design wind load [knm/m2] h height above the ground [m] qpok linear load in the experiment [kn/m] m moment on the handlebar section [knm/m] sf safety factor [/] references [1] akter s. t., khani m. s. characterisation of laminated glass for structural applications. master thesis, 2013 [2] “saint gobain”. updated mar. 2015, accessed 13 apr. 2015. http://uk.saint-gobain-glass.com [3] neugebauer j. design of glass structures, scientific symposium: future trends in civil engineering, zagreb, croatia, 2014. [4] “manual for design of guardrails”. accessed 05 jun. 2015. http://www.allium.com/pdf/manuel-en.pdf [5] larcher m., solomos g., casadei f., gebbeken n. experimental and numerical investigations of laminated glass subjected to blast loading. international journal of impact engineering 2012; 39:42-5 doi:10.1016/j.ijimpeng.2011.09.006 [6] müller r., wagner m. berechnung sprengwirkungshemmender fenster und fassadenkonstruktionen. bauingenieur 2008;81(11):475-87 [7] sun d., andrieux f., ockewitz a., klamser h., hogenmüller j. modelling of the failure behaviour of windscreens and component tests. 5th european lsdyna users conference, 25-26 may, 2005. [8] morison c., zobec m., frenceschet a. the measurement of pvb properties at high strain rates, and their application in the design of laminated glass under bomb blast. isiems 2007, international symposium on interaction of the effects of munitions with structures, 17-21 september,orlando, us, 2007. [9] bennison s.j., jagota a., smith c.a. fracture of glass/poly(vinyl butyral)(butacite)laminates in biaxial flexure. journal of the american ceramic society 1999; 82(7):1761-70. [10] duser a.v., jagota a., bennison s.j. analysis of glass/polyvinyl butral laminates subjected to uniform pressure. journal of engineering mechanics 1999; 125(4):435-42. doi:10.1061/(asce)0733-9399(1999)125:4(435) [11] rak m., damjanović d., krolo j.: report on testing the glass beam, universitiy of zagreb, faculty of ce., 2013. [12] “sap 2000” 16 eval, computers and structures, inc., 2013. 70 http://dx.doi.org/10.1016/j.ijimpeng.2011.09.006 http://dx.doi.org/10.1061/(asce)0733-9399(1999)125:4(435) acta polytechnica ctu proceedings 3:65–70, 2016 1 introduction 2 bridge project details 2.1 bridge construction 2.2 railing element 2.3 wind load 3 load bearing capacity test 4 simulation of load bearing capacity test in sap 2000 5 results 6 conclusion list of symbols references 234 acta polytechnica ctu proceedings 2(1): 234–237, 2015 234 doi: 10.14311/app.2015.02.0234 the nova v5584 sgr: a short review r. poggiani 1 1department of physics, università di pisa corresponding author: rosa.poggiani@df.unipi.it abstract the nova v5584 sgr was discovered during 2009 october. it has been monitored in different domains of the electromagnetic spectrum: optical, infrared and x-rays. the optical and infrared observations suggest that v5584 sgr is a fe ii nova that formed dust. no x-ray emission was observed around the time of maximum. keywords: cataclysmic variables -novae optical spectroscopy individual: v5584 sgr. 1 introduction nova sagittarii no. 4 was discovered by nishiyama and kabashima (2009) on 2009 october 26 and later designated v5584 sgr (samus, 2009). v5584 sgr is a classical example of multi-wavelength astrophysics, since it has been observed in different parts of the electromagnetic spectrum: optical, infrared and x-rays. the very early optical spectroscopic observations secured by kinugasa et al. (2009), maehara et al. (2009), fujii (maehara et al., 2009), munari et al. (2009) in the optical and by raj et al. (2009) in the infrared showed that v5584 sg is a fe ii nova, in the context of the classification by williams et al. (1991), williams (1992). spectra secured by russell et al. (2010) during 2010 february showed that dust formation had occurred. poggiani (2011) has monitored v5584 sgr during the late decline. v5584 has been investigated in the optical domain by the stony brook/smarts consortium (walters et al., 2012). the present paper reviews the history of v5584 sgr observations. 2 light curve the photometric evolution and the main parameters of v5584 sgr has been reported by poggiani (2011) and will be briefly summarized below for completeness. the v band light curve is reported in fig. 1. the epoch of maximum is mjd=55134.208 (2009 october 29), while the decline time by two magnitudes is 27±2 days, making v5584 sgr a moderately fast nova, according to the classification by payne-gaposchkin (1957). the reddening of v5584 sgr is 0.82±0.12. the estimated distance of v5584 sgr is in the interval 5.8-7.1 kpc. the absolute magnitude at maximum is in the range -7.2....7.7, while the white dwarf mass is in the range 0.8-0.9 m�. all parameters extracted by the analysis of the light curve suggest that v5584 sgr is a fe ii nova, according to the classification by della valle & livio (1998). the spectroscopic observations described below will provide further evidence for the classification. 55100 55150 55200 55250 55300 55350 55400 55450 55500 time (mjd) 8 9 10 11 12 13 14 15 16 17 v m a g n it u d e figure 1: v band photometry of v5584 sgr 3 optical spectroscopy spectroscopic observations of v5584 sgr around maximum have been secured by different authors. kinugasa et al. (2009) observed hα and fe ii lines in emission with p cyg profiles on 2009 october 27. the presence of hα in emission was confirmed by maehara et al. (2009) on the same day. munari et al. (2009) observed a highly reddened absorption continuum on october 28, with hα and fe ii showing an emission component and the same continuum with faint emission from balmer ad fe ii multiplets on october 29. it was suggested that v5584 sgr was a fe ii nova caught around maximum, according to the classification by williams et al. (1991), williams (1992). v5584 sgr has also been observed at the higashi-hiroshima observatory with the 234 http://dx.doi.org/10.14311/app.2015.02.0234 the nova v5584 sgr: a short review kanata telescope (uemura, 2009) on october 31 and november 3, 5 1. the intensity of emission lines progressively strengthened, while the continuum decayed. v5584 sgr is part of the stony brook smarts spectral atlas of southern novae that includes data of more than sixty novae (walters et al., 2012). their observations of v5584 sgr on november 11 showed balmer lines, fe ii multiplets and o i. poggiani (2011) has observed v5584 sgr during the late decline (after the seasonal gap) at the loiano observatory, italy, using the 152 cm telescope and the bfosc imager and spectrograph. the observation of june 4 is reported in fig. 2. the spectrum has been secured using grism #4 that has a range 4000 to 8500 å and a resolution of 3.97 å/pixel, reduced with bias and flat frames, extracted using the optimal extraction method by horne (1986), calibrated in wavelength with an hear lamp and corrected with the instrumental response. the spectrum shows a prominent hα transition, intense forbidden lines of [o iii], [n ii] and the high ionization line [fe vii] 6087, typical of a nova in the nebular stage a0 (williams et al., 1991; williams 1992). the same pattern of lines has been observed by the smarts consortium on october 17. figure 2: optical spectrum of v5584 sgr during the nebular stage (poggiani, 2011) the line profiles are a signature of the ejecta distribution. optically thin winds and optically thin shells show rounded and rectangular line profiles, respectively. the profile of hα at two epochs (labelled with the number of days elapsed from maximum) is reported in fig. 3 (poggiani (2011)). the profiles show a red wing, probably explained by the emerging of [n ii] 6584. the rest wavelength is coincident with the 0 km/s label. the observations of v5584 sgr during the nebular stage are part of a monitoring program (poggiani, 2012) active since 2005 at the loiano observatory, italy, using the 152 cm telescope and the bfosc imager and spectrograph. the most part of the spectra are secured with grism #4 to build an homogeneous data set of spectra over the range 4000 to 8500 å. the monitoring aims to build a spectral atlas and includes most novae observable at the telescope latitude that underwent outburst after 2005. other sources are v1663 aql, v1722 aql, v2362 cyg, v2467 cyg, v2468 cyg, v2491 cyg, v407 cyg, kt eri, v959 mon, v2670 oph, v496 sct, v5558 sgr, v458 vul, v459 vul. the novae are observed at several epochs, from the early decline to the nebular stage and on the way to quiescence. figure 3: line profiles of hα during the nebular stage (poggiani, 2011) the importance of monitoring of novae over long time intervals has been addressed by several contributors at this conference (pagnotta, 2013), (ederoclite, 2013), (tappert, 2013). at the moment of writing v5584 sgr is very far from quiescence. at the moment of discovery, corelli (2009) showed that nothing was visible at the nova position on a palomar plate with limiting magnitude 21. 4 infrared observations the relevance of observations of novae in the infrared has been discussed by banerjee and ashok (2012) and by chesneau and banerjee (2012). banerjee and ashok (2012) suggest that the fe ii and he/n classes share a set of common spectral lines in the near infrared, namely hydrogen, helium, nitrogen and oxygen. on the other hand, there is a clear signature belonging only to novae of fe ii class, the presence of strong carbon lines: thus the near infrared observations can provide an additional suggestion of the spectroscopic class of a nova. several novae, mostly novae belonging to fe ii class, produce dust during the decline. the dust presence is marked by a rise of infrared emission and can become apparent also as a dip in the optical light curve. the last signature is not necessarily observed, since dust can be arranged as an optically thin shell. 1http://f.hatena.ne.jp/kanataobslog/20091106162703 235 r. poggiani infrared observations of v5584 sgr have been secured by raj et al. (2009) on october 29 at mt. abu telescope showed paschen, brackett, o i, c i, n i lines with p cyg profiles, whose components wew separated by about 550-650 km/s. the emission components strengthened, while the absorption components faded by november 3. the transitions observed in spectra are typical of the fe ii class novae, providing an independent confirmation of the classification of v5584 sgr. the following observations in the infrared were secured after the seasonal gap. on 2010 february 10 russell et al. (2010) observed v5584 sgr in the region 3-14 µm with aeos telescope, discovering that the nova had formed dust since the previous observations. the infrared continuum was dominated by the thermal emission of the dust at a temperature of 880±50 k; no details of the dust composition or the derivation of the temperature are reported. the above infrared observatons provide a clue that was not available due to the lack of observations during the seasonal gap. 5 x-ray observations the importance of x-ray observations of novae has been addressed by several authors, also at this conference (ness, 2013), (orio, 2013). v5584 sgr has been observed with the monitor of all-sky x-ray image (maxi) (shimanoe et al., 2010). the system is on board of iss and uses two high sensitivity x-ray detectors, the gas slit camera (32 ccd chips) and the solid state slit camera (a xenon filled proportional conunter). maxi scans all sky every 92 minutes. shimanoe et al. (2010) have searched the prompt x-ray emission at the ignition of the thermonuclear runaway. the authors suggested that classical novae could emit x-rays at the outburst in analogy to the type i x-ray bursts of x-ray binaries. the observations included several peculiar novae: v1723 aql, v407 cyg, v2673 and v2674 oph, v1722 aql, kt eri, v496 sct. the authors have focused on the archive data in the energy band 1.5-4 kev, the lowest energy available, assuming that nova outbursts preferentially emit soft x-rays. they also investigated the 4-10 kev and 10-20 kev bands. no prompt emission was detected in the three bands for any nova. there are no reported x-ray observations of v5584 sgr at later stages, when supersoft emission could have occurred. 6 conclusions v5584 sgr is a standard fe ii nova. the sinergy of observations in different domains allowed a classification of the spectral class and the discovery of dust production during the seasonal gap. acknowledgement the author thanks the organizers for the invitation to this very stimulating workshop. thanks to the tac committee of the loiano observatoy for the allotted time and to i. bruni for his precious assistance. the author is grateful to the anonymous referee for the careful comments. references [1] banerjee, d. p. k. and ashok, n. m.: 2012, basi 40, 243. [2] chesneau, o. and banerjee, d. p. k.: 2012, basi 40, 267. [3] corelli, p.: 2009, cbet 1994, 1. [4] della valle, m. & livio, m.: 1998, apj 506, 818. doi:10.1086/306275 [5] ederoclite, a.: these proceedings [6] horne, k.: 1986, pasp 98, 609. doi:10.1086/131801 [7] kinugasa, k. et al.: 2009, iauc 9089, 2. [8] maheara, h. et al.: 2009, iauc 9089, 3. [9] munari, u. et al.: 2009, cbet 1999, 1. [10] ness, j. u.: these proceedings [11] nishiyama, k. & kabashima, f.: 2009, iauc 9089, 1. [12] orio, m.: these proceedings [13] pagnotta, a.: these proceedings [14] payne-gaposchkin, c.: 1957, the galactic novae, north holland, amsterdam. [15] poggiani, r.: 2011, astrophy. space sci. 333, 115. doi:10.1007/s10509-011-0595-z [16] poggiani, r.: 2012, mem. sait 83, 753 [17] raj, a. et al., 2009: cbet 2002, 1. [18] russell, r. w. et al.: 2010, iauc 9118, 2. [19] samus, n. & kazarovets, e. v.: 2009, iauc 9089, 4. [20] shimanoe, j. et al.: 2010, the first year of maxi: monitoring variable x-ray sources, 4th international maxi workshop held november 30 december 2, tokyo, japan [21] tappert, k.: these proceedings 236 http://dx.doi.org/10.1086/306275 http://dx.doi.org/10.1086/131801 http://dx.doi.org/10.1007/s10509-011-0595-z the nova v5584 sgr: a short review [22] uemura, m. on behalf of the kanata team: 2009, asp conf. ser. 404, 69 [23] walters, f. m. et al.: 2012, pasp 124, 1057. doi:10.1086/668404 [24] williams, r. e. et al.: 1991, apj 376, 721. doi:10.1086/170319 [25] williams, r. e.: 1992, aj 104, 725. discussion martin henze: you described x-ray observations and mentioned soft band of 1.5-4 kev in which soft x-rays would be expected. i just wanted to comment that in this band we would not expect to see supersoft x-ray emission which normally does not go above 1 kev. rosa poggiani: the maxi astronomers stated that they were looking for soft x-ray emission in the above energy range. i agree that to detect the x-ray emission you are mentioning you should use another instrument jan-uwe ness: the maxi collaboration was looking for prompt emission. 237 http://dx.doi.org/10.1086/668404 http://dx.doi.org/10.1086/170319 introduction light curve optical spectroscopy infrared observations x-ray observations conclusions acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0043 acta polytechnica ctu proceedings 4:43–49, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app assessment of burnable absorber fuel design by uwb1 depletion code martin loveckýa, ∗, jana jiřičkováa, radek škodaa, b a regional innovation centre for electrical engineering, university of west bohemia, univerzitní 8, 306 14 plzeň, czech republic b faculty of mechanical engineering, czech technical university, technická 4, 160 07 praha 6, czech republic ∗ corresponding author: lovecky@rice.zcu.cz abstract. uwb1 depletion code is being developed as a fast computational tool for the study of burnable absorbers in university of west bohemia in pilsen, czech republic. the research of fuel depletion aims at a development and introduction of advanced types of burnable absorbers in nuclear fuel. burnable absorbers compensate for the initial excess reactivity and consequently allow for lower power peaking factors and longer fuel cycles with higher fuel enrichments. the paper describes the depletion calculations of candu, pwr and sfr nuclear fuel doped with rare earth oxides as burnable absorber. uniform distribution of burnable absorber in the fuel is assumed. based on performed depletion calculations, rare earth oxides are divided into two groups, suitable burnable absorbers and poisoning absorbers. moreover, basic economic comparison is performed based on actual stock prices. keywords: burnable absorber, fuel depletion, monte carlo. 1. introduction uwb1 nuclear fuel depletion code [1, 4] is being developed as a fast fuel depletion code to conduct burnable absorber research. the goal of the research is to optimize new materials as burnable absorbers (ba) in nuclear fuel. bas compensate for the initial excess reactivity and consequently allow for lower power peaking factors and longer fuel cycles with higher fuel enrichments. research of advanced types of burnable absorbers (ba) in nuclear fuel requires fast depletion code that would be able to calculate broad range of elements, nuclides or their combination. state-of-art depletion codes require large amount of computational time, therefore, decision to develop fast depletion code was made. currently developed uwb1 depletion code comprises of transport and burnup solver that works in 2spc depletion scheme in order to perform fast calculation of nuclear fuel depletion. in order to achieve high precision in uwb1 code, monte carlo solver for lattice cell calculations is used to solve transport equation. bateman equations are solved in special depletion scheme that relies on nuclide-based predictorcorrector method in both transport and burnup part of fuel depletion. the paper describes the depletion calculations of candu, pwr and sfr nuclear fuel doped with rare earth oxides as burnable absorber. uniform distribution of burnable absorber in the fuel is assumed. based on performed depletion calculations, rare earth oxides are divided into two groups, suitable burnable absorbers and poisoning absorbers. moreover, basic economic comparison is performed based on current stock prices. 2. uwb1 depletion code first version of the newly developed uwb1 fast nuclear fuel depletion code [1] significantly reduced calculation time by omitting the solution step for the boltzmann transport equation. however, estimation of multiplication factor during depletion was not sufficiently calculated [2]. moreover, 1-group effective cross sections for strong absorber models like gadolinium showed disagreement between the uwb1 tested code and the serpent reference code [3]. monte carlo transport solver for uwb1 code was introduced [4] in order to improve code accuracy and remove pre-calculated case-dependent data libraries, therefore, eliminate constant effective cross section assumption. nuclear data from endf/b-vii.1 library is used for uwb1 monte carlo solver. standard raytracing algorithm for neutron random walk is a part of the solver. two-dimensional fuel pin models geometry, where fuel lattice model is assumed. the speed of the solver is higher than for the mcnp6 reference code. for light water reactor models, uwb1 monte carlo solver is on average 10 times faster. two-step predictor-corrector method (2spc) was developed for uwb1 code. the idea is to change the coupling of transport and burnup solvers by omitting major fraction of transport solver callings, because transport solver is, by orders of magnitude, slower than burnup solver. both transport and burnup variables are calculated for predicted states and corrected with more precise values as the two parts of fuel depletion are coupled. only three transport solver solutions are used, the initial fuel state and predicted and corrected states for final burnup state. effective cross sections are evaluated during fuel depletion by 43 http://dx.doi.org/10.14311/ap.2016.4.0043 http://ojs.cvut.cz/ojs/index.php/app m. lovecký, j. jiřičková, r. škoda acta polytechnica ctu proceedings assuming nuclide-based non-linear dependency. multiplication factors in the depletion steps other than the first and the last one are estimated by neutron production to absorption ratio that is calculated without the need to call the transport solver. employing 2spc depletion scheme leads to a significant reduction of calculation time by a factor of 10. with the monte carlo speed-up against the reference code, fuel depletion with uwb1 code is around 100 times faster than with mcnp6 reference code. 3. burnable absorbers in order to mark a material as a good burnable absorber, two properties are desired. these are a high neutron absorption cross section with low daughter nuclide’s neutron absorption cross section. high neutron capture cross section causes neutrons to be absorbed, therefore, multiplication factor of the fuel is decreased and initial excess of reactivity is compensated. the absorber is burnable if the nuclide resulting from neutron absorption has a lower neutron absorption cross section, therefore, reactivity worth of inserting burnable absorber is decreasing during fuel depletion. the higher the differences between cross sections, the faster the burnable absorber will be burned. typical absorption reaction for burnable absorbers is (n,γ), although another reactions are possible, e.g. (n,α) for light elements like boron. present nuclear fuels use mainly gadolinium, europium and erbium oxides. gadolinium oxide is burned faster than the two others and because of the very high absorption cross section of gd-157 nuclide, initial reactivity compensation for gadolinium higher than for other materials. gadolinium nuclides with interesting burnable properties are gd-155 and gd-157 with natural abundances 14.80 at% and 15.65 at%. europium has natural abundance of 47.81 at% eu-151 and 52.19 at% eu-153. both europium nuclides behaves as a good burnable absorber nuclide, both nuclides have have high neutron capture cross sections with end-product nuclide characterized by low capture cross section. erbium nuclides with interesting burnable properties are er-166 and er-167 with natural abundances 33.50 at% and 22.87 at%. total cross section for thermal energy 0.0253 ev are: 60 801 barns for gd-155, 253 929 barns for gd-157, 9190 barns for eu-151, 367 barns for eu153, 31 barns for er-166 and 652 barns for er-167. 4. rare earth oxides burnable absorber research will be conducted in two phases. the first phase will be represented by parametric study of various types of ba materials. the uwb1 depletion code will be used in order to calculate a large number of cases, employing its fast calculation scheme. promising materials will be used in the second phase for a deeper neutronic analysis with state-of-art codes, possible discrepancies or inaccuracies of the uwb1 code will be removed. the output of the research is to sort materials by their suitability as ba from neutronics point of view. after that, another types of analyses, like chemical compatibility, thermal conductivity, economic evaluation and other relevant areas will be performed. the paper focuses only on rare earth oxides as burnable absorbers. these are often used or planned to use as ba in pwr nuclear fuel. the most common is gadolinium oxide, followed by europium and erbium oxides. rare earth oxides were briefly analyzed in [5] for vver nuclear fuel, in this paper, pwr, candu and sfr nuclear fuels are compared. rare earth elements are available as metals or oxides. thermal reactors and some of proposed fast reactors use uranium oxide fuel, therefore, rare earth oxides were evaluated. seventeen rare earth elements are abundant in nature (scandium, yttrium and 15 lanthanides from lanthanum to lutetium). most of them are available in oxides with x2o3 stechiometry, in order to introduce the same normalization for all elements, masses of all other rare earth cases were recalculated to have x2o3 stechiometry. 5. calculation cases candu, pwr and sfr nuclear fuel depletion was calculated. of seventeen available rare earth oxides, 16 were calculated. promethium, having only radioactive nuclides, was excluded from the study. it the first part of the study, criticality calculations with uwb1 code were performed in order to determine rare earth oxides content in the fuel (wt% x2o3). the target multiplication factor was selected to be around 0.05 lower than for the fresh fuel without ba, namely 1.10 for candu fuel, 1.25 for pwr fuel and 1.15 for sfr fuel (fresh fuel multiplication factors are 1.15, 1.32, 1.20). in the second part, fuel depletion with uwb1 code were calculated and multiplication factor dependency on fuel burnup was analyzed. finally, economic comparison was performed. uwb1 code requires simple fuel cell geometry. recently, the option to include arbitrary number of concentric cylinders with either square or triangular lattice was included. candu fuel was homogenized and modeled as 18 concentric cylinders that represents four fuel rings of the total number of 37 pins. natural uranium oxide fuel with final burnup 10 000 mwd/mtu was assumed. fuel depletion was divided into 31 steps with constant power 32.4364 mw/mtu. standard candu materials were used in the model – zircaloy-4 for fuel cladding, zr2.5nb alloy for pressure tube that is divided by co2 gap from zircaloy-2 calandria tube, heavy water with density 0.8179 g/cm3 and purity 99.11 wt% as coolant in the pressure tube and heavy water with density 1.086 99 g/cm3 and purity 99.97 wt% as moderator between calandria tube. pwr nuclear fuel was modeled as 5.0 wt% u-235 enriched uranium oxide with zircaloy-4 cladding and light water moderator with 600 ppm boric acid. fuel radius 0.4025 cm and cladding radius 0.4750 cm was 44 vol. 4/2016 assessment of burnable absorber fuel design by uwb1 depletion code used, lattice half pitch 0.6295 cm was assumed as typical pwr dimensions. the fuel was depleted with power 40.0 mw/mtu in 34 steps up to final burnup 50 000 mwd/mtu. sfr nuclear fuel was modeled as 15.0 wt% u-235 enriched uranium oxide with 7.9 g/cm3 iron cladding and 0.88 g/cm3 sodium coolant. fuel radius 0.28 cm and cladding radius 0.33 cm was used, lattice half pitch 0.42 cm were used as typical sfr dimensions (bn-800 reactor was assumed as typical sfr reactor). the fuel was depleted with power 130.0 mw/mtu in 74 steps up to final burnup 150 000 mwd/mtu, three times higher than of pwr fuel. 6. criticality results candu, pwr and sfr nuclear fuel criticality calculation results from uwb1 code are presented in fig. 1 to fig. 3. the figures show dependency of multiplication factor for fresh fuel with various ba content (mass fraction) on logarithm scale. 1 e 0 5 1 e 0 4 1 e 0 3 1 e 0 2 1 e 0 1 1 e + 0 0 0 . 0 0 . 2 0 . 4 0 . 6 0 . 8 1 . 0 1 . 2 mu ltip lic ati on fa cto r ( -) b a w e i g h t c o n t e n t ( ) 2 1 s c 3 9 y 5 7 l a 5 8 c e 5 9 p r 6 0 n d 6 2 s m 6 3 e u 6 4 g d 6 5 t b 6 6 d y 6 7 h o 6 8 e r 6 9 t m 7 0 y b 7 1 l u c a n d u : m u l t i p l i c a t i o n f a c t o r v s b a c o n t e n t figure 1. candu multiplication factor vs ba content. 1 e 0 5 1 e 0 4 1 e 0 3 1 e 0 2 1 e 0 1 1 e + 0 0 0 . 0 0 . 2 0 . 4 0 . 6 0 . 8 1 . 0 1 . 2 1 . 4 mu ltip lic ati on fa cto r ( -) b a w e i g h t c o n t e n t ( ) 2 1 s c 3 9 y 5 7 l a 5 8 c e 5 9 p r 6 0 n d 6 2 s m 6 3 e u 6 4 g d 6 5 t b 6 6 d y 6 7 h o 6 8 e r 6 9 t m 7 0 y b 7 1 l u p w r : m u l t i p l i c a t i o n f a c t o r v s b a c o n t e n t figure 2. pwr multiplication factor vs ba content. multiplication factor reduction from ba insertion into the fuel is higher for candu fuel that needs only a small ba content to achieve target multiplication 1 e 0 5 1 e 0 4 1 e 0 3 1 e 0 2 1 e 0 1 1 e + 0 0 0 . 2 0 . 4 0 . 6 0 . 8 1 . 0 1 . 2 mu ltip lic ati on fa cto r ( -) b a w e i g h t c o n t e n t ( ) 2 1 s c 3 9 y 5 7 l a 5 8 c e 5 9 p r 6 0 n d 6 2 s m 6 3 e u 6 4 g d 6 5 t b 6 6 d y 6 7 h o 6 8 e r 6 9 t m 7 0 y b 7 1 l u s f r : m u l t i p l i c a t i o n f a c t o r v s b a c o n t e n t figure 3. sfr multiplication factor vs ba content. factor. pwr fuel is similar to candu, however, due to higher fuel enrichment, more ba content is needed. lastly, sfr with fast neutron spectra and high fuel enrichment requires considerably higher ba content than previous fuel types in order to achieve desirable initial reactivity compensation. based on performed criticality calculations and selected target multiplication factor values 1.10 for candu fuel, 1.25 for pwr fuel and 1.15 for sfr fuel, ba weight fraction was calculated. linear interpolation between selected ba contents (3 values for each logarithm order) was used. calculated ba content is shown in fig. 4 to fig. 6. 2 1 s c 3 9 y 5 7 l a 5 8 c e 5 9 p r 6 0 n d 6 2 s m 6 3 e u 6 4 g d 6 5 t b 6 6 d y 6 7 h o 6 8 e r 6 9 t m 7 0 y b 7 1 l u 1 e 0 6 1 e 0 5 1 e 0 4 1 e 0 3 1 e 0 2 1 e 0 1 1 e + 0 0 ba ox ide w eig ht fra cti on () b u r n a b l e a b s o r b e r t y p e ( ) c a n d u : b u r n a b l e a b s o r b e r w e i g h t f r a c t i o n figure 4. candu burnable absorber weight fraction. in the case of both candu and pwr fuel, even maximum considered ba content (50 wt% x2o3) was not enough for dropping the multiplication factor to the target one. for pwr fuel, 58-ce content was also too high. for sfr fuel, it was possible to achieve desired target multiplication factor for all rare earth oxides ba. 45 m. lovecký, j. jiřičková, r. škoda acta polytechnica ctu proceedings 2 1 s c 3 9 y 5 7 l a 5 8 c e 5 9 p r 6 0 n d 6 2 s m 6 3 e u 6 4 g d 6 5 t b 6 6 d y 6 7 h o 6 8 e r 6 9 t m 7 0 y b 7 1 l u 1 e 0 5 1 e 0 4 1 e 0 3 1 e 0 2 1 e 0 1 1 e + 0 0 ba ox ide w eig ht fra cti on () b u r n a b l e a b s o r b e r t y p e ( ) p w r : b u r n a b l e a b s o r b e r w e i g h t f r a c t i o n figure 5. pwr burnable absorber weight fraction. 2 1 s c 3 9 y 5 7 l a 5 8 c e 5 9 p r 6 0 n d 6 2 s m 6 3 e u 6 4 g d 6 5 t b 6 6 d y 6 7 h o 6 8 e r 6 9 t m 7 0 y b 7 1 l u 1 e 0 3 1 e 0 2 1 e 0 1 1 e + 0 0 ba ox ide w eig ht fra cti on () b u r n a b l e a b s o r b e r t y p e ( ) s f r : b u r n a b l e a b s o r b e r w e i g h t f r a c t i o n figure 6. sfr burnable absorber weight fraction. minimum ba content was calculated for gadolinium oxide for candu and pwr fuel (7.52 × 10−6 and 1.02 × 10−4 weight fraction) and for europium oxide for sfr fuel (8.67 × 10−3 weight fraction). for most of rare earth oxides, candu and pwr fuel requires ba content from 0.1 wt% x2o3 to 1.0 wt% x2o3. for sfr fuel, required ba content mostly varies from 1.0 wt% x2o3 to 10.0 wt% x2o3, i.e. 10 times higher than of thermal spectra fuels. maximum ba content was calculated for 58-ce. for candu fuel, 26.8 wt% ce2o3 was determined. slightly lower conten of 20.4 wt% ce2o3 was calculated for sfr fuel. pwr fuel required more than 50.0 wt% ce2o3 and cerium oxide was excluded, making 39-y the case of maximum ba content for pwr fuel (40.5 wt% y2o3). 7. depletion results candu, pwr and sfr nuclear fuel criticality calculation results from uwb1 code are presented in fig. 7 to fig. 9. 0 2 0 0 0 4 0 0 0 6 0 0 0 8 0 0 0 1 0 0 0 0 0 . 9 0 0 . 9 5 1 . 0 0 1 . 0 5 1 . 1 0 1 . 1 5 mu ltip lic ati on fa cto r ( -) b u r n u p ( m w d / m t u ) n o _ b a 6 0 n d 6 7 h o 2 1 s c 6 2 s m 6 8 e r 3 9 y 6 3 e u 6 9 t m 5 7 l a 6 4 g d 7 0 y b 5 8 c e 6 5 t b 7 1 l u 5 9 p r 6 6 d y c a n d u : f u e l d e p l e t i o n w i t h b a figure 7. candu multiplication factor during fuel depletion. candu figure shows that from macroscopic view, only 63-eu and 64-gd compensate initial reactivity in the way that initial reactivity transient due to fission product build-up almost disappears. for other ba, initial reactivity is compensated, however, initial reactivity transient is still presented. for good burnable absorbers, residual poisoning (decrease of multiplication factor compared to no ba case) is negligible, other ba have residual poisoning that, for some cases, is even higher than initial reactivity compensation. 0 1 0 0 0 0 2 0 0 0 0 3 0 0 0 0 4 0 0 0 0 5 0 0 0 0 0 . 8 5 0 . 9 0 0 . 9 5 1 . 0 0 1 . 0 5 1 . 1 0 1 . 1 5 1 . 2 0 1 . 2 5 1 . 3 0 1 . 3 5 mu ltip lic ati on fa cto r ( -) b u r n u p ( m w d / m t u ) n o _ b a 6 0 n d 6 7 h o 2 1 s c 6 2 s m 6 8 e r 3 9 y 6 3 e u 6 9 t m 5 7 l a 6 4 g d 7 0 y b 5 8 c e 6 5 t b 7 1 l u 5 9 p r 6 6 d y p w r : f u e l d e p l e t i o n w i t h b a figure 8. pwr multiplication factor during fuel depletion. for pwr fuel, similar conclusions as for candu fuel can be deduced. only 63-eu and 64-gd oxides are able to change the shape of multiplication factor dependency on burnup, i.e. remove initial reactivity transient. however, for pwr fuel, initial reactivity transient is not only removed, but the effect is reversed, 46 vol. 4/2016 assessment of burnable absorber fuel design by uwb1 depletion code multiplication factor increases in the first days of fuel operation. for sfr fuel, all rare earth oxides show similar behavior, burning of the oxides is very slow and reactivity compensation remains relatively constant, therefore, rare earth oxides behaves mostly as non-burnable absorbers. 0 2 5 0 0 0 5 0 0 0 0 7 5 0 0 0 1 0 0 0 0 0 1 2 5 0 0 0 1 5 0 0 0 0 0 . 9 0 0 . 9 5 1 . 0 0 1 . 0 5 1 . 1 0 1 . 1 5 1 . 2 0 mu ltip lic ati on fa cto r ( -) b u r n u p ( m w d / m t u ) n o _ b a 6 0 n d 6 7 h o 2 1 s c 6 2 s m 6 8 e r 3 9 y 6 3 e u 6 9 t m 5 7 l a 6 4 g d 7 0 y b 5 8 c e 6 5 t b 7 1 l u 5 9 p r 6 6 d y s f r : f u e l d e p l e t i o n w i t h b a figure 9. sfr multiplication factor during fuel depletion. based on performed depletion calculations and multiplication factor progress during fuel depletion, ba worth was calculated. ba worth is defined as reactivity differences between two cases, the one with ba in the fuel and the one without ba. results are depicted in fig. 10 to fig. 12. 0 2 0 0 0 4 0 0 0 6 0 0 0 8 0 0 0 1 0 0 0 0 8 0 0 0 6 0 0 0 4 0 0 0 2 0 0 0 0 ba w ort h ( pc m) b u r n u p ( m w d / m t u ) n o _ b a 6 0 n d 6 7 h o 2 1 s c 6 2 s m 6 8 e r 3 9 y 6 3 e u 6 9 t m 5 7 l a 6 4 g d 7 0 y b 5 8 c e 6 5 t b 7 1 l u 5 9 p r 6 6 d y c a n d u : b a w o r t h figure 10. candu burnable absorber worth during fuel depletion. from ba worth curve variation during fuel depletion, it can be stated which rare earth oxides behaves like more or less suitable burnable absorbers and which behaves like poisoning absorbers. the latter group can be divided into constant poisoning absorbers (ba worth is constant function of burnup) and increasingly poisoning absorbers (ba worth is decrasing function of decrasing function of burnupburnup, i.e. diverges 0 1 0 0 0 0 2 0 0 0 0 3 0 0 0 0 4 0 0 0 0 5 0 0 0 0 1 2 0 0 0 1 0 0 0 0 8 0 0 0 6 0 0 0 4 0 0 0 2 0 0 0 0 ba w ort h ( pc m) b u r n u p ( m w d / m t u ) n o _ b a 6 0 n d 6 7 h o 2 1 s c 6 2 s m 6 8 e r 3 9 y 6 3 e u 6 9 t m 5 7 l a 6 4 g d 7 0 y b 5 8 c e 6 5 t b 7 1 l u 5 9 p r 6 6 d y p w r : b a w o r t h figure 11. pwr burnable absorber worth during fuel depletion. 0 1 0 0 0 0 2 0 0 0 0 3 0 0 0 0 4 0 0 0 0 5 0 0 0 0 5 0 0 0 4 0 0 0 3 0 0 0 2 0 0 0 1 0 0 0 0 ba w ort h ( pc m) b u r n u p ( m w d / m t u ) n o _ b a 6 0 n d 6 7 h o 2 1 s c 6 2 s m 6 8 e r 3 9 y 6 3 e u 6 9 t m 5 7 l a 6 4 g d 7 0 y b 5 8 c e 6 5 t b 7 1 l u 5 9 p r 6 6 d y s f r : b a w o r t h figure 12. sfr burnable absorber worth during fuel depletion. from zero reactivity worth, rather than converging behaviour of burnable absorbers). because of the fact that ba worth curve does not change its derivation sign (constant, increasing or decreasing function), it is possible to define ratio of eoc to boc reactivity worth and compare it to unity, see final results in fig. 13. candu and pwr fuel behaves similarly, burnable absorbers are more depleted for final burnup in the case of candu fuel, therefore, residual poisoning is lower for candu and burnable absorbers are more suited for candu fuel. total number of 9 out of 16 rare earth oxides are marked as burnable absorbers for both candu and pwr fuels, namely 21-sc, 60-nd, 62-sm, 63-eu, 64gd, 66-dy, 67-ho, 68-er and 69-tm. total number of 4 out of 16 rare earth oxides are marked as burnable absorbers only for candu fuel, but behaves as poisoning absorbers for pwr fuel. these are 39-y, 57-la, 59-pr and 71-lu. the rest 3 out of 16 rare earth oxides, 65-tb, 58-ce and 70-yb, are poisoning absorbers for both candu and pwr fuel. 47 m. lovecký, j. jiřičková, r. škoda acta polytechnica ctu proceedings 2 1 s c 3 9 y 5 7 l a 5 9 p r 6 0 n d 6 2 s m 6 3 e u 6 4 g d 6 5 t b 6 6 d y 6 7 h o 6 8 e r 6 9 t m 7 1 l u 0 . 0 0 . 5 1 . 0 1 . 5 2 . 0 ba w ort h e oc / b a wo rth b oc () b a e l e m e n t ( ) c a n d u p w r s f r b a w o r t h r a t i o b e t w e e n e o c a n d b o c figure 13. rare earth oxides as ba eoc/boc worth ratio. situation for sfr fuel is simpler than for both thermal spectra fuels. only 2 out of 16 rare earth oxides, 66-dy and 69-tm, are suitable as burnable absorbers. thulium oxide is a far better ba for sfr than dysprosium oxide that burns very slowly. however, significant fraction of thulium oxide remains in the fuel at eoc. 8. economic comparison economic comparison of rare earth oxides as burnable absorbers is based on prices from early 2016 [6], however, rare earth oxides are mined mainly in one country and prices are volatile. table 1 (or fig. 14) summarizes the results. input oxides prices (rmb/mt) were converted to match x2o3 stechiometry and multiplied by target ba content (weight fraction) to evaluate oxides prices for unit uranium mass (rmb/mtu). total fuel price comprises of ba price, uranium price and other factors, only ba price is compared in this study. element price (rmb/mt) 39-y 34 000 57-la 13 250 58-ce 13 900 59-pr 429 000 60-nd 322 500 62-sm 17 000 63-eu 1900 64-gd 87 500 65-tb 4300 66-dy 1915 68-er 265 000 table 1. rare earth oxides as ba – economic comparison. 3 9 y 5 7 l a 5 9 p r 6 0 n d 6 2 s m 6 3 e u 6 4 g d 6 5 t b 6 6 d y 6 8 e r 1 e 0 1 1 e + 0 0 1 e + 0 1 1 e + 0 2 1 e + 0 3 1 e + 0 4 1 e + 0 5 ba pr ice (r mb /m tu ) b a e l e m e n t ( ) c a n d u p w r s f r r a r e e a r t h o x i d e p r i c e s a s b u r n a b l e a b s o r b e r s figure 14. rare earth oxides as ba – economic comparison. for the best burnable absorbers, the economic comparison showed the price of rare earth oxide in mtu vary from 1 rmb to 1000 rmb for candu and pwr fuels and approximately 100 times higher prices for sfr fuels. 9. conclusions uwb1 depletion code was used to assess burnable absorber design in the nuclear fuel. candu, pwr and sfr fuel lattice criticality and depletion calculations were performed for rare earth oxides as burnable absorbers uniformly distributed in the fuel. rare earth oxides content needed to achieve target initial reactivity compensation was evaluated. based on multiplication factor behavior during fuel depletion and ba reactivity worth ratio for final burnup and inital state, rare earth oxides were divided into two groups, burnable absorbers and poisoning absorbers. candu and pwr fuel behaves similarly. total number of 9 out of 16 rare earth oxides are marked as burnable absorbers for both candu and pwr fuels, namely 21-sc, 60-nd, 62-sm, 63-eu, 64-gd, 66-dy, 67ho, 68-er and 69-tm. total number of 4 out of 16 rare earth oxides are marked as burnable absorbers only for candu fuel, but behaves as poisoning absorbers for pwr fuel. these are 39-y, 57-la, 59-pr and 71-lu. the rest 3 out of 16 rare earth oxides, 65-tb, 58-ce and 70-yb, are poisoning absorbers for both candu and pwr fuel. situation for sfr fuel is simpler than for both thermal spectra fuels. only 2 out of 16 rare earth oxides, 66-dy and 69-tm, are suitable as burnable absorbers. thulium oxide is a far better ba for sfr than dysprosium oxide that burns very slowly. for the best burnable absorbers, the economic comparison showed the price of rare earth oxide in mtu 48 vol. 4/2016 assessment of burnable absorber fuel design by uwb1 depletion code vary from 1 rmb to 1000 rmb for candu and pwr fuels and approximately 100 times higher prices for sfr fuels. further analysis of possible burnable absorbers with broad range of elements, nuclides or their combination is in progress. the research aims to a quantitative comparison of materials that could be used as burnable absorbers. the resulting data are planned to be utilized in subsequent fuel design analyses including thermal conductivity, chemical compatibility, irradiation stability, economic evaluation etc. list of symbols boc beginning of cycle (irradiation) eoc end of cycle (irradiation) mt metric ton mtu metric tons of uranium rmb renminbi (chinese currency system) acknowledgements research and development has been funded by tacr, project no. te01020455 – centre for advanced nuclear technologies (canut). references [1] m. lovecky, l. piterka, j. prehradny, r. skoda. uwb1 – fast nuclear fuel depletion code. annals of nuclear energy 71:333–339, 2014. doi:10.1016/j.anucene.2014.04.021. [2] j. prehradny, m. lovecky, r. skoda. burnable absorber comparison between vver, pwr and sfr with uwb1 and serpent codes. in proceedings of the 2014 22nd international conference on nuclear engineering. 2014. doi:10.1115/icone22-30894. [3] j. leppanen. development of a new monte carlo reactor physics bode. ph.d. thesis, helsinki university of technology, 2007. [4] m. lovecky, j. jirickova, r. skoda. monte carlo solver for uwb1 nuclear fuel depletion code. annals of nuclear energy 85:778–7879, 2015. doi:10.1016/j.anucene.2015.06.035. [5] j. prehradny, m. lovecky, r. skoda. rare earth oxides as burnable absorber for vver nuclear fuel. in proceedings of the 2015 23rd international conference on nuclear engineering. 2015. [6] shanghai metals market, prices of rare earth oxide, http://www.metal.com/metals/rare-earth/prices. 49 http://dx.doi.org/10.1016/j.anucene.2014.04.021 http://dx.doi.org/10.1115/icone22-30894 http://dx.doi.org/10.1016/j.anucene.2015.06.035 http://www.metal.com/metals/rare-earth/prices acta polytechnica ctu proceedings 4:43–49, 2016 1 introduction 2 uwb1 depletion code 3 burnable absorbers 4 rare earth oxides 5 calculation cases 6 criticality results 7 depletion results 8 economic comparison 9 conclusions list of symbols acknowledgements references 66 acta polytechnica ctu proceedings 1(1): 66–70, 2014 66 doi: 10.14311/app.2014.01.0066 evolution of massive population iii stars sung-chul yoon1 1astronomy program, department of physics and astronomy, seoul national university, seoul, 151-742, republic of korea corresponding author: yoon@astro.snu.ac.kr abstract while the evolution of massive stars in the local universe is dominated by mass-loss, the evolution of massive population iii stars should be dominated by rotation. an important effect of rotation is rotationally-induced chemical mixing that can dramatically change the stellar structure and the nucleosynthesis. this has significant consequences in the predicted explosion types of population iii stars. keywords: stars: evolution stars: rotation stars: population iii supernovae gamma-ray bursts. 1 introduction formation of population iii (pop iii) stars marks the end of the so-called dark age in the early universe. these first stars are believed to be intrinsically massive (30 m� < m < 1000 m�), given that the primordial gas does not contain any efficient coolants in star-forming regions, and supposed to play a key role in the evolution of the early universe, in many aspects (see bromm 2013, for a recent review). neutral hydrogens and heliums are re-ionized by the first lights from these first generations of stars, and the primordial gas becomes polluted with heavy elements produced by both hydrostatic and explosive nucleosyntheses in them. although direct identification of them will be highly unlikely even with next generations of telescopes, their nature can be indirectly probed with observations of their explosions that may be luminous enough to be discovered in near future. observational studies on the surface abundances of extremely metal poor stars, which are believed to retain the nucleosynthesis signatures of pop iii stars, will also provide a good constraint on the evolution of pop iii stars. theoretical predictions that can confront future observations are indispensable for these efforts (e.g., marigo et al. 2001; heger & woosley 2002; umeda & nomoto 2003; heger & woosley 2010; kasen et al. 2011; limongi & chieffi 2012; dessart et al. 2013; whalen et al. 2013; tanaka et al. 2013). in this paper, i discuss the current understandings on the evolution of massive pop iii stars, focusing on the role of rotation in the stellar evolution (e.g., marigo et al. 2003; ekström et al. 2008; yoon et al. 2012). recent numerical studies indicate that massive pop iii stars would be born with surface rotation close to the critical value (e.g., stacy et al. 2011). 2 importance of rotation in our galaxy, stars having masses higher than about 30 m� usually undergo strong mass-loss during their evolution. this is because numerous metal lines in their atmospheres and the consequent high radiation pressure cause instabilities of various sorts at the stellar surfaces (see puls et al. 2008, for a recent review). the evolution of massive stars at high metallicity is therefore dominated by mass-loss. for instance, in the local universe, stars as massive as 120 m� at their birth would lose more than 100 m� throughout their lives, dying as 10 20 m� stars (e.g., meynet & maeder 2005). this also means that such massive stars would rapidly lose most of their initial angular momentum, rendering the effect of rotation minor or negligible in general. by contrast, it is believed that massive pop iii stars do not experience significant mass-loss because they are generally very stable (baraffe et al. 2001). in addition, lines by hydrogen and helium are too weak to drive radiation-driven winds (krticka & kubát, 2006). therefore, massive pop iii stars would retain a significant fraction of the initial angular momentum throughout their evolution. this has the following important consequence: while the evolution of very massive stars at high metallicity is dominated by mass-loss, the evolution of massive pop iii stars is dominated by rotation. an important effect of rotation is chemical mixing resulting from rotationally-induced hydrodynamic instabilities like eddington-sweet circulations that are driven by thermal imbalance inside rapidly rotating stars (e.g., maeder 2009). see brott et al. (2011b) for a 66 http://dx.doi.org/10.14311/app.2014.01.0066 evolution of massive population iii stars ! ! !" !" !"#$%&'()"&*+,"./0$,1%&&2'3"$"40-0"*5'()"&*+,"!"#$ #%&'($)*+)',) -"#$ #$ %& '(% )* )+% , -*.+/'(%)*)+%, !0/(%1",' 2,3"$%." 45"'6%("'+7' "88+6+",)$0' 9(*:"/'/%&,' /;"')%' )5"'6%("<",3"$%."' 6%;.$+,1= 65*%&'1"#071"&&%850 >%'6%("'9(*:+,1'90' )5"'6%("<",3"$%."' 6%;.$+,1= (98&"5,"-'8":0#0;' <2'#"+%+,"-'&,=0'>?@ #)(%,1'65"?+6*$'1(*/+",)'9")&"",' )5"'!<9;(,+,1'6%,3"6)+3"'6%("' *,/')5"'",3"$%." >%'65"?+6*$'1(*/+",)'9")&"",' )5"'!<9;(,+,1'6%,3"6)+3"'6%("' *,/')5"'",3"$%." figure 1: illustration for the bifurcation of massive star evolution according to their initial rotational velocity. slowly rotating massive stars develop the classical core-envelope structure such that they become red-supergiant stars during the post-main sequence phases. rapidly rotating massive stars, on the other hand, may undergo very rapid rotationally-induced chemical mixing to such an extent that quasi-chemical homogeneity can be maintained on the main sequence, if their initial metallicity is sufficiently low. in this case, stars are gradually transformed into helium stars at the end of the main sequence. see yoon & langer (2005), yoon et al. (2006), and woosley & heger (2006) for more details. recent dicussion on the observational tests and related uncertainties of the rotational mixing scenario. the hydrogen burning core is supplied with fresh hydrogenrich material from the envelope because of this mixing, resulting in a larger core size than in the corresponding case without rotation. the most extreme case is the socalled chemically homogeneous evolution (che), which is illustrated in fig. 2. with a sufficiently high rotational velocity (i.e., more than about 50% of the local keplerian value at the equatorial surface), the chemical mixing timescale by eddington-sweet circulations can be shorter than the nuclear burning time of hydrogen, so as to maintain quasi-chemical homogeneity in the star. because almost all the hydrogens in the star participate nuclear burning in this way, the star gradually becomes a massive helium star at the end of the main sequence. this is an important evolutionary channel to make massive he stars at low metallicity, while che may not play an important role at high metallicity because of rapid angular momentum loss via stellar winds (see yoon et al. 2006). this mode of evolution is also considered an important channel to produce long gamma-ray bursts from metal-poor massive stars (yoon & langer 2005; yoon et al. 2006; woosley & heger 2006). stars with the che evolve bluewards instead of evolving redwards on the hr diagram. the presence of such stars in the early universe would make significant impact on the history of reionization, because they produce more ionizing photons by several factors than stars that follow the normal evolution (yoon et al. 2012). this particular mode of stellar evolution also results in a significant diversity of pop iii star explosions, as discussed below. 3 explosions of massive population iii stars the theoretical predictions on the final fates of massive pop iii stars are given as the phase diagram in fig. 2. pop iii stars that develop the classical core-envelope structure would die as red or blue supergiant stars, and their final fates may not be much different from those 67 sung-chul yoon 10 20 35 60 100 160 250 500 0.0 0.2 0.4 0.6 0.8 final fates of rotating massive pop iii stars 10 20 35 60 100 160 250 mzams [msun] 0.0 0.2 0.4 0.6 0.8 v z a m s /v k forbidden region v zams =v crit sn iip (ns remnant) collapse to bh, or weak sn ii collapse to bh pisn (sn ii) p u ls. p is n (s n ii) grb / hn (sn ibc) p u ls. p is n g r b /h n (ib c) pisn (sn ibc) figure 2: final fates of massive pop iii stars for given initial masses and rotational velocities, predicted from the stellar evolution models by yoon et al. (2012). the rotational velocity on the zero-age main sequence is given in units of the keplerian value at the equatorial surface (i.e., vk = √ gm/r). we adopted the most up-to-date calibration for the rotationally-induced chemical mixing efficiency by brott et al. (2011a). the region for the chemically homogeneous evolution (che) is marked by yellow color, and the thick solid line gives the boundary between the che regime and the region for the normal evolution. different regions are marked by expected final outcomes: type ii supernovae (sn ii), collapse to black hole (bh), pulsational pair-instability supernovae (puls-pisn), pair-instability supernovae (pisn), long gamma-ray bursts (grb), hyper-novae (hn), type ib or ic supernovae (sn ibc). in the forbidden region, the equatorial surface velocity exceeds the critical velocity: vcrit = √ (gm/r)[1 − γ], where γ is the eddington factor at the stellar surface. this figure is taken from yoon et al. (2012). of non-rotating pop iii stars as predicted by heger & woosely (2002). but the limits of the initial masses for pulsational pair-instability and pair instability supernovae would move downwards for a higher initial rotational velocity, because rotational mixing tends to increase the he core masses. as explained above, in the che regime, pop iii stars would die as massive helium stars and therefore produce supernovae of type ib/c. conditions for long gamma-ray bursts can also be fulfilled in this regime for an initial mass range of 13 84 m�. while non-rotating pop iii stars would require initial masses of about 140 270 m� to produce a pair instability supernova, this mass range would decrease to about 84 190 m� in the che regime (see also chatzopoulos & wheeler 2012). another interesting prediction is that for intial masses of about 56 84 m�, a pulsational pair-instability supernova would be followed by a long gamma-ray bursts, only several years later. given that gamma-ray jets would interact with a very massive circumstellar matter in this case, extremely bright afterglow might be produced. our models also predict that the final masses of pop iii grb progenitors (∼ 10 − 70 m�) would be significantly more massive than those of grb progenitors in the local universe (∼ 10 − 20 m�; yoon et al. 2006). on the other hand, recently several authors argued for super-collapsar formation in very massive pop iii stars (m > 270 m�; mészáros & rees 2010; komissarov & barkov 2010; suwa & ioka 2011 ). however, our models show that very massive pop iii stars with initial masses larger than about 190 m� cannot retain enough angular momentum in the core to produce any rotation-powered event at their deaths. this is because such stars have a very large convective core that leads to efficient angular momentum transport from the innermost layers to the stellar surface, and because large eddington-factors of these stars facilitates mass and angular momentum loss via centrifugally driven massshedding. 68 evolution of massive population iii stars acknowledgement i am very grateful to norbert langer and alexandra dierks for their contribution to this work. references [1] baraffe, i., heger, a., & woosley, s.e., 2001, apj, 550, 890 doi:10.1086/319808 [2] bromm, v., rep. prog. phys., 2013, submitted. [3] brott, i., de mink, s.e., & cantiello, m. et al., 2011a, a&a, 530, 115 [4] brott, i., evans, c.j., & hunter, i. et al., 2011b, a&a, 530, 115 [5] chatzopoulos, e., & wheeler, j.c., 2012, apj, 760, 154 doi:10.1088/0004-637x/760/2/154 [6] dessart, l., waldman, r., livne, e., hillier, d.j., blindin, s., 2013, mnras, 428, 3227 doi:10.1093/mnras/sts269 [7] ekström, s., meynet, g., chiappini, c., hirschi, r., & maeder, a., 2008, a&a, 489, 685 [8] heger, a., & woosley, s.e., 2002, apj, 567, 532 doi:10.1086/338487 [9] heger, a., & woosley, s.e., 2010, apj, 724, 341 doi:10.1088/0004-637x/724/1/341 [10] kasen, d., woosley, s.e., & heger, a., 2011, apj, 734, 102 doi:10.1088/0004-637x/734/2/102 [11] komissarov, s.s., & barkov, m.v., 2010, mnras, 402, 25 doi:10.1111/j.1745-3933.2009.00792.x [12] krticka, j. & kubát, j., 2006, a&a, 446, 1039 [13] limongi, m., & chieffi, a., 2012, apjs, 199, 38 doi:10.1088/0067-0049/199/2/38 [14] maeder, a., 2009, physics, formation and evolution of rotating stars, astronomy and astrophysics labirary, springer [15] marigo, p., chiosi, c., & kudritzki, r.-p., 2003, a&a, 399, 617 [16] marigo, p., girardi, l., chiosi, c., & wood, p.r., 2001, a&a, 371, 152 [17] mészáros, p., & rees, m.j., 2010, apj, 715, 967 doi:10.1088/0004-637x/715/2/967 [18] meynet, g., & maeder, a., 2005, a&a, 429, 581 [19] puls, j., vink, j.s., & najarro, f., 2008, a&arv, 16, 209 [20] spruit, h., 2002, a&a, 381, 923 [21] stacy, a., bromm, v., & loeb, a., 2011, mnras, 413, 543 doi:10.1111/j.1365-2966.2010.18152.x [22] suijs, m.p.l., langer, n., poelarends, a.-j., yoon, s.-c., hger, a., & herwig, f., 2008, a&a, 481, 87 [23] suwa, y., & ioka, k., 2011, apj, 726, 107 doi:10.1088/0004-637x/726/2/107 [24] tanaka, m., moriya, t.j., & yoshida, n., 2013, arxiv:1306.3743 [25] umeda, h., & nomoto, k., 2003, nature, 422, 871 doi:10.1038/nature01571 [26] whalen, d. et al., 2013, apj, 762, 6 doi:10.1088/2041-8205/762/1/l6 [27] woosley, s.e., & heger, a., 2006, apj, 637, 914 doi:10.1086/498500 [28] yoon, s.-c., & langer, n., 2005, a&a, 443, 643 [29] yoon, s.-c., langer, n., & norman, c., 2006, a&a, 460, 199 [30] yoon, s.-c., dierks, a., & langer, n., 2012, a&a, 542, 113 discussion dorota rosinska: you have shown that evolution of low-metallicity massive stars strongly depends on rotation. such stars should rotate differentially. what laws of differential rotation you assume? how results depend on the degree of differential rotation (ωe/ωc; ωe is the angular velocity at the equatorial surface, and ωc the angular velocity at the center)? sung-chul yoon: stars always tend to rotate differentially because of the contraction of the core that results in spinning-up the central layers. the degree of differential rotation in a star is then determined by the efficiency of the angular momentum transport from the core to the envelope. i.e., we do not ”assume” the degree of differential rotation, but it is self-consistently calculated with a given prescription for the transport of angular momentum. in our models, we adopted the socalled talyer-spruit dynamo (spruit 2002) which leads to a very strong coupling between the core and the envelope, making a star rotate almost rigidly on the main sequence. stellar evolution models including the taylerspruit dynamo have proved to be consistent with observations, in terms of the predicted spin-rates of stellar 69 http://dx.doi.org/10.1086/319808 http://dx.doi.org/10.1088/0004-637x/760/2/154 http://dx.doi.org/10.1093/mnras/sts269 http://dx.doi.org/10.1086/338487 http://dx.doi.org/10.1088/0004-637x/724/1/341 http://dx.doi.org/10.1088/0004-637x/734/2/102 http://dx.doi.org/10.1111/j.1745-3933.2009.00792.x http://dx.doi.org/10.1088/0067-0049/199/2/38 http://dx.doi.org/10.1088/0004-637x/715/2/967 http://dx.doi.org/10.1111/j.1365-2966.2010.18152.x http://dx.doi.org/10.1088/0004-637x/726/2/107 http://dx.doi.org/10.1038/nature01571 http://dx.doi.org/10.1088/2041-8205/762/1/l6 http://dx.doi.org/10.1086/498500 sung-chul yoon remnants like white dwarfs and neutron stars (suijs et al. 2008). matteo guainazzi: which black hole spin distribution would you expect from your evolutionary scenario? sung-chul yoon: bimodal distribution. if a bh forms in the che regime, its kerr parameter will be close to 1. otherwise, bhs will not have any significant angular momentum, having a kerr parameter close to zero. 70 introduction importance of rotation explosions of massive population iii stars 194 acta polytechnica ctu proceedings 1(1): 194–199, 2014 194 doi: 10.14311/app.2014.01.0194 multi-wavelength view of supernova remnants manami sasaki1 1institute for astronomy and astrophysics, eberhard karls university tübingen, sand 1, d-72076 tübingen, germany corresponding author: sasaki@astro.uni-tuebingen.de abstract this contribution gives a very short overview on the emission of supernova remnants and the processes that are responsible for both the thermal and non-thermal origins of the emission, typically observed in radio, x-rays, and up to γ-rays. we discuss in particular the case of the galactic snr ctb 109. as detailed x-ray studies combined with observations in radio have shown, ctb 109 is interacting with a giant molecular cloud complex. the interaction of the snr shock with dense interstellar clouds is responsible for both the unusual semi-circular morphology of the snr and the bright x-ray feature inside the snr, and, as has been shown recently, seems also to play a major role in the production of γ-rays. keywords: supernova remnants radio ir optical x-rays γ-rays. 1 introduction supernova remnants (snrs) are objects that are formed by a supernova (sn) explosion at the end of the life of a star. owing to the shock waves of the sn explosion, the stellar ejecta as well as the ambient medium is heated to temperatures higher than a million kelvin. therefore, snrs are responsible for the heating of the interstellar medium (ism) and the distribution of heavier elements that were processed inside a star. in addition, both electrons and nuclei can be accelerated in the shock waves of the snrs to relativistic energies. therefore, snrs are also considered to be one source of galactic cosmic rays. by now, about 280 snrs are known in our galaxy and about 50 in the magellanic clouds. detailed studies of snrs in the galaxy and the magellanic clouds allow us to understand the nature of their shocks, the various radiation processes that are responsible for the complex emission observed over the full electromagnetic spectrum from radio to γ-rays, the interaction of the shock with the cooler, and often inhomogeneous ism, and the acceleration of particles. 2 snrs at lower energies 2.1 radio emission in the strong shock waves of snrs, electrons are accelerated and interact with the interstellar magnetic fields. these electrons emit synchrotron radiation and thus make the snr bright in radio. the distribution of the surface brightness of the electron synchrotron emission in snrs can be obtained from radio imaging observations. in general, the radio emission consists of narrow shell-like or filamentary emission and additional diffuse emission. the reason for these two kinds of synchrotron emission is that particles with higher energies lose their energy faster through synchrotron emission than those with lower energies. the long-lived lower-energy particles can diffuse further away from the shock and radiate over a much wider area, producing the more extended diffuse emission with a steeper spectrum. figure 1: radio continuum emission from snr da 530 (1420 mhz) in contour representation (left) and gray scale (right). vectors in the left image show the polarized intensity by their length and the direction of the e-field by their orientation (landecker et al., 1999). since synchrotron emission is polarized, radio observations can reveal the magnetic field structure in and around snrs. observations have shown that young remnants in general have a radial field structure. for remnants in the free expansion phase and in the early phases of the adiabatic expansion, radial motions due 194 http://dx.doi.org/10.14311/app.2014.01.0194 multi-wavelength view of supernova remnants to rayleigh-taylor instabilities at the contact discontinuity between the shocked ejecta and the shocked ism can enhance the radial magnetic fields. on the contrary, older remnants have a more tangential field structure with higher polarized fractions (milne, 1987). the polarized fraction in younger remnants are typically p = 10 − 15%, while older remnants like snr da 530 can have a polarized fraction of p = 50% (fig. 1, landecker et al., 1999). 3 thermal x-ray emission the outward moving shock wave of an snr heats the interstellar medium, while the reverse shock that propagates inward heats the ejecta of the sn explosion. as a whole the ionized gas inside an snr has temperatures of ∼ 106−7 k typically. in plasmas the excitation and ionization of the atoms are mainly induced by electrons that collide with ions. snrs have very low densities and thus are optically thin. they can be considered to be out of thermal and ionization equilibrium, especially when they are still young. if the time after the plasma has been shocked is short, only a few collisions have occurred that cause ionization. the ionization states of the ions at a given temperature are thus lower than in the case of collisional ionization equilibrium (cie). instead, the plasma in snrs is characterized by non-equilibrium ionization (nei). the shape of the continuum of the thermal plasma is primarily determined by the electron temperature, while the temperature of ions might be different than that of electrons. the line emission in the x-ray band can be used to derive the ionization states of the ions as well as the element abundances. therefore, the x-ray spectrum can, on the one hand, reveal the element distribution in the ejecta, and on the other hand, the typical abundances of the ism in which the snr expands. 3.1 charge exchange newer observations of snrs in x-rays have revealed spectral features indicating charge exchange (cx) between the hot thin plasma of an snr and the cold interstellar gas. x-ray emission from cx can in principle be produced at any astrophysical site where hot plasma interacts with (partially) neutral gas, and was originally discoverd in x-ray data of comets (lisse et al., 1996, cravens, 1997). first observational evidences of cx in snrs had been obtained as a broad component of h α emission (e.g., chevalier et al., 1980). observations of the rim of the cygnus loop using suzaku and xmmnewton have revealed a significant emission feature at 0.7 kev, which was identified to be a complex of he-like o k(γ + δ + �) and was interpreted as cx emission from interaction between he-like o and neutrals (fig. 2, katsuda et al., 2011). figure 2: non-x-ray background subtracted spectrum of a northeastern region of the cygnus loop snr taken with suzaku along with the best-fit model consisting of a bremsstrahlung plus gaussians (katsuda et al., 2011). line identifications are given on top of the upper panel. the gaussian shown in red represent the cx emission. the x-ray background components are shown in green. 3.2 overionization new suzaku observations of mixed-morphology (mm) snrs in the last few years have shown that in many cases the thermal x-ray emission can be interpreted as that from an overionized, recombining plasma in ionization non-equilibrium (e.g., yamaguchi et al., 2009, see fig. 3). mixed-morphology snrs, also called thermalcomposite remnants, are usually found in dense regions near molecular clouds and show very little or no shell emission. the x-ray emission is bright in the (projected) interior and is of thermal nature. the signature of an overionized plasma is a radiative recombination continuum (rrc) and strong ly emission of, e.g., si, s, or fe. in general, snrs are underionized and the shocked plasma slowly develops to ionization equilibrium. therefore, it is rather surprising to observe overion195 manami sasaki ization in snrs. to reach overionization in a plasma above ∼ 106 k, electrons have to cool rapidly. possible processes for such a cooling are either fast expansion, thermal conduction, or radiative cooling with enhanced abundances, which are all processes that are likely to occur in particular in snrs evolving in significantly inhomogenous ism. figure 3: suzaku xis spectrum in the 1.75 – 6.0 kev band (black: fi, red: bi). the best fit model is shown including the cxb (black) and vapec and gaussians for the source emission. rrc emisison of h-like mg, si, and s are shown in magenta (yamaguchi et al., 2009). 4 snrs as particle accelerators 4.1 non-thermal x-ray emission supernova remnants are non-thermal, extended radio sources with a flux spectrum described by a power-law with sν ∝ να with α ≈ −0.5. the synchrotron radiation is emitted by a non-thermal population of relativistic electrons, which can be described by a powerlaw. synchrotron emission in radio indicates that there are electrons with energies in the gev range in the snr shell. synchrotron emission from an snr in the x-ray band was confirmed for the first time for the remnant of sn 1006 (koyama et al., 1995) based on asca data, which showed that the non-thermal emission is dominating the emission of two outer shell regions. soon after, more snrs were also confirmed to radiate non-thermal x-ray emission: cas a (e.g., the et al., 1996, allen et al., 1997), rx j1713.7-3946 (g347.30.5, pfeffermann & aschenbach, 1996), rx j0852.04622 (g266.2-1.2, vela jr, aschenbach 1998). the xray emission of rx j1713.7-3946 and rx j0852.0-4622 is completely of non-thermal nature with no thermal emission detected so far. x-ray synchrotron emission is only observed in younger snrs and is indicative of electrons with energies of up to 1014 ev. the angular resolution of the newer x-ray telescopes, in particular that of the chandra x-ray observatory, made it possible to reveal narrow filaments of non-thermal x-ray emission close to the outer shock in several young snrs: sn 1006, rx j1713.7-3946, rx j0852.0-4622, cas a, tycho, and kepler. these filaments are very thin (∼ 1 − 2′′) and indicate relatively high magnetic fields (b = 100 − 600µg). recently, a deep chandra observation of tycho’s snr revealed bright non-thermal stripes, clearly seen in the energy band of > 4kev (fig. 4, eriksen et al., 2011). the stripes are likely caused by peaks in magnetic fields. figure 4: chandra image of the tycho’s snr in the energy band of 4.0 – 6.0 kev. the boxes show the regions with stripes or filaments of non-thermal emission (eriksen et al., 2011). 4.2 gamma-ray emission in the last decade some snrs have been detected with γ-ray observatories and are now well established astrophysical γ-sources. there are two major origins of γ-ray emission in snrs: • leptonic: interactions of relativistic electrons with low-energy photons of the cosmic microwave background cause inverse compton up-scattering of the photons. in addition, the non-thermal electrons will also emit bremsstrahlung through interactions with ions in the snr. • hadronic: accelerated nuclei can produce γray emission if they collide with atomic nuclei, thereby creating, among others, neutral π0s, which decay into γ-ray photons. this is believed to be enhanced in snrs near high-density clouds where the probability for the collision of the accelerated protons from the snr with cold protons or nuclei is higher (e.g., ic 443, w44, see fig. 5). the gev and tev emission of snrs can thus give us a direct view of the accelerated particles. 196 multi-wavelength view of supernova remnants figure 5: gamma-ray spectra of the snrs ic 443 and w44 and fitted models (ackermann et al., 2013). 5 the galactic snr ctb 109 the galactic supernova remnant ctb 109 (g109.1– 1.0) is the host of the anomalous x-ray pulsar (axp) 1e 2259+586 (fahlman & gregory, 1981) and has an interesting semi-circular morphology in both the x-ray and the radio. it is believed to be located next to a giant molecular cloud (gmc) complex and is thus one of the most striking examples of an interaction of an snr with a molecular cloud. the snr was discovered in x-rays with einstein by gregory & fahlman (1980). hughes et al. (1981) identified the radio counterpart in the galactic plane survey at 49 cm with the westerbork synthesis radio telescope. its shell is incomplete on the western side both in x-rays and in radio. the semi-circular radio and x-ray morphology of ctb 109 suggests the shock has been stopped by the gmc complex in the west. an elongated feature in co (‘co arm’) extends from the gmc complex to an x-ray minimum inside the snr in the north (tatematsu et al., 1987), indicating that a part of the dense gas of the gmc complex extends in front of the remnant. we performed observations of the entire snr ctb 109 with xmm-newton using the european photon imaging cameras (epics). a mosaic of all observations is shown in fig. 6. as mentioned before, ctb 109 seems to have expanded into a dense medium in the west, while the shock wave is expanding into a lowerdensity medium in the east. the radii of the northern and the southern periphery of ctb 109 are slightly smaller than to the east suggesting that ctb 109 is confined by denser material also in the north and the south. however, the relative difference is only ∼5 – 10%. therefore, the overall shape of the shell can be considered as a semi-sphere. there is an x-ray bright interior region known as the lobe, which is brighter than any part of the snr shell. the x-ray spectrum from the lobe obtained with xmm-newton is completely thermal (sasaki et al., 2004). we found co clouds around the lobe in new high-resolution co data from the five college radio observatory (sasaki et al., 2006). one of the co clouds overlaps with the lobe and has an additional component towards higher negative velocities in the velocity profile. we believe that the bright x-ray emission of the lobe is the result of the interaction of the snr shock with a dense molecular cloud, which was hit by the blast wave on the western side. the main part of the cloud was heated and is now visible as the lobe, while some denser parts might have survived. figure 6: mosaic rgb image of the snr ctb 109 created from the xmm-newton observations: red = 0.3 – 0.9 kev, green = 0.9 – 1.5 kev, blue = 1.5 – 4.0 kev (sasaki et al., 2004). 197 manami sasaki 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 vnei+vnei fe (0.0 5.0 x solar) 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 vnei+vnei abund. (0.0 5.0 x solar) 1 1.2 1.4 1.6 1.8 2 2.2 2.4 2.6 2.8 3 vnei+vnei si significance (1 3 ) 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 vnei+vnei si (0.0 5.0 x solar) 1 1.2 1.4 1.6 1.8 2 2.2 2.4 2.6 2.8 3 vnei+vnei fe significance (1 3 ) (b) (d) (e) (c)(a) figure 7: best fit abundances obtained for the emission of snr ctb 109 from the analysis of the chandra acis spectra by assuming a thermal model that consists of two vnei components for thermal non-equilibrium plasma for the shocked ism and for the shocked ejecta (sasaki et al., 2013). the shown parameters are: abundances for elements other than si, s, and fe fitted for the ejecta emission (a), si abundance fitted for the ejecta emission (b) and its significance calculated as (abundance of the si – abundance of the other elements)/error of the abundance of si (c), fe abundance and its significance (d and e, respectively). through the analysis of a deep, high-resolution data of the region around the lobe taken with chandra advanced ccd imaging spectrometer (acis), we discovered ejecta emission inside the snr west of the lobe (sasaki et al., 2013). the distriubtion of the siabundance and fe-abundance are shown in in fig. 7. these images show the parameter values obtained from the analysis of the spectra extracted in the regions indicated. the abundances measured for the ejecta emission are all comparable to or lower than solar values, except for the lobe. here, the si and fe abundances are in particular significantly higher. the enhanced abundances in the lobe suggest that there is an ejecta clump or a conglomeration of an ejecta clump and a shocked cloud. at the position of ctb 109 a gev source was detected with fermi (castro et al., 2012, see fig. 8). this γ-ray source is located southwest of the lobe where the snr appears dark in x-rays and thus is most likely highly absorbed. castro et al. (2012) performed a comprehensive modeling of the emission of ctb 109, in which they could reproduce the fluxes in radio, x-rays, and γ, as well as the x-ray spectrum taken with the xmm-newton epic in a model including a significant contribution by π0-decay. this is again consistent with the picture that there is significant interaction between the snr shock and a dense interstellar cloud. 6 summary being driven by strong shock waves formed in supernova explosions, supernova remnants are one of the main drivers of the matter cycle and dynamic evolution of galaxies. they are responsible for the chem198 multi-wavelength view of supernova remnants ical enrichment of galaxies. the shock waves create new structures in the ism by forming a cavity filled with hot thin gas and sweep the colder gas around it. particles are accelerated in the shock waves and their emission is observed over the entire range of the observable electromagnetic spectrum. in recent years observations of x-ray synchrotron radiation confirmed the existence of electrons with energies in the tev range in snr shocks. newest gev and tev obervations indicate inverse compton scattering as well as pion-decay processes in and around snr shocks. self-consistent modeling of the spectrum from radio to tev as in the case of the galactic snr ctb 109 helps us to understand both the heating and acceleration processes in snr shocks. figure 8: smoothed fermi-lat count maps in the range 2 – 200 gev (castro et al., 2012). white contours show the x-ray emission (0.5 – 5.0 kev) from xmm-newton data. acknowledgement m.s. acknowledges support by the deutsche forschungsgemeinschaft through the emmy noether research grant sa 2131/1. references [1] ackermann, m., et al. 2013, science, 339, 807 doi:10.1126/science.1231160 [2] allen, g. e., et al. 1997, apjl, 487, 97 [3] aschenbach, b. 1998, nature 396, 141 doi:10.1038/24103 [4] castro, d., slane, p., ellison, d. c., patnaude, d. j. 2012, apj, 756, 88 doi:10.1088/0004-637x/756/1/88 [5] chevalier, r. a., kirshner, r. p., raymond, j. c. 1980, apj, 235, 186 [6] cravens, t. e. 1997, geophys. res. lett., 24, 105 [7] eriksen, k. a., et al. 2011, apjl, 728, 28 doi:10.1088/2041-8205/728/2/l28 [8] fahlman, g. g. & gregory, p. c. 1981, nature, 293, 202 doi:10.1038/293202a0 [9] gregory, p. c. & fahlman, g. g. 1980, nature, 287, 805 doi:10.1038/287805a0 [10] hughes, v. a., harten, r. h., costain, c. h., van den bergh, s. 1981, apj, 246, l127 doi:10.1086/183568 [11] [12] katsuda, s., et al. 2011, apj, 730, 24 doi:10.1088/0004-637x/730/1/24 [13] koyama, k., et al. 1995, nature, 378, 255 doi:10.1038/378255a0 [14] landecker, t. l., et al. 1999, apj, 527, 866 [15] lisse, c. m., et al. 1996, science, 274, 205 doi:10.1126/science.274.5285.205 [16] milne, d. k., 1987, aujph, 40, 771 [17] pfeffermann, e., aschenbach, b. 1996, proc. of ‘röntgenstrahlung from the universe’, 267 [18] sasaki, m., kothes, r., plucinsky, p. p., gaetz, t. j., brunt, c. m. 2006, apj, 642, l149 doi:10.1086/504844 [19] sasaki, m., et al. 2004, apj, 617, 322 [20] sasaki, m., plucinsky, p. p., gaetz, t. j., bocchino, f. 2013, a&a, 552, 45 [21] tatematsu, k., et al. 1987, a&a, 184, 279 [22] the, l.-s., et al. 1996, a&as, 120, 357 [23] yamaguchi, h., et al. 2009, apjl, 705, 6 doi:10.1088/0004-637x/705/1/l6 discussion carlotta pittori’s comment: to distinguish between ic and π0 decay in the gev range, at least in some cases like w44, the new planck data combined with agile and fermi data are very much constraining. the mw picture in the case of middle-aged snrs interacting with molecular clouds seem to largely favour the hadronic scenario (π0 decay). 199 http://dx.doi.org/10.1126/science.1231160 http://dx.doi.org/10.1038/24103 http://dx.doi.org/10.1088/0004-637x/756/1/88 http://dx.doi.org/10.1088/2041-8205/728/2/l28 http://dx.doi.org/10.1038/293202a0 http://dx.doi.org/10.1038/287805a0 http://dx.doi.org/10.1086/183568 http://dx.doi.org/10.1088/0004-637x/730/1/24 http://dx.doi.org/10.1038/378255a0 http://dx.doi.org/10.1126/science.274.5285.205 http://dx.doi.org/10.1086/504844 http://dx.doi.org/10.1088/0004-637x/705/1/l6 introduction snrs at lower energies radio emission thermal x-ray emission charge exchange overionization snrs as particle accelerators non-thermal x-ray emission gamma-ray emission the galactic snr ctb 109 summary acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0008 acta polytechnica ctu proceedings 4:8–12, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app post critical heat transfer and fuel cladding oxidation vojtěch caha∗, jakub krejčí department of nuclear reactors, faculty of nuclear sciences and physical engineering, czech technical university in prague, czech republic ∗ corresponding author: vojtech.caha@fjfi.cvut.cz abstract. the knowledge of heat transfer coefficient in the post critical heat flux region in nuclear reactor safety is very important. although the nuclear reactors normally operate at conditions where critical heat flux (chf) is not reached, accidents where dryout occur are possible. most serious postulated accidents are a loss of coolant accident or reactivity initiated accident which can lead to chf or post chf conditions and possible disruption of core integrity. moreover, this is also influenced by an oxide layer on the cladding surface. the paper deals with the study of mathematical models and correlations used for heat transfer calculation, especially in post dryout region, and fuel cladding oxidation kinetics of currently operated nuclear reactors. the study is focused on increasing of accuracy and reliability of safety limit calculations (e.g. dnbr or fuel cladding temperature). the paper presents coupled code which was developed for the solution of forced convection flow in heated channel and oxidation of fuel cladding. the code is capable of calculating temperature distribution in the coolant, cladding and fuel and also the thickness of an oxide layer. keywords: heat transfer, post dryout, cladding oxidation, isolated channel. 1. introduction the post critical heat flux heat transfer is encountered when the surface temperature becomes too high to maintain a continuous liquid contact, and the surface becomes covered by a continuous vapour blanket. it results in coverage of the heated surface by a continuous vapour film in the case of film boiling regime, or an intermittent vapour film in the case of transition boiling regime. the boundary between these post dryout heat transfer boiling regimes is the minimum of film boiling temperature, or leidenfrost temperature. the typical boiling curve is depicted in figure 1. post dryout heat transfer is initiated as soon as the critical heat flux condition is reached and it remains until rewetting or quenching of the surface takes place. the post critical heat flux heat transfer regimes in boiling flow can be divided as: (1.) transition boiling regime, (2.) flow film boiling regime, (a) inverted annular film boiling, (b) slug flow film boiling, (c) dispersed flow film boiling. the transition boiling is a combination of unstable film boiling and unstable nucleate boiling consecutively existing at a given location on a heating surface. whereas, film boiling is generally defined as a boiling where only the vapour phase is in contact with the heated surface. inverted annular film boiling is characterized by a vapour layer separating the continuous liquid core from the heated surface and usually encountered at void fractions below 40 %. while dispersed flow film boiling is characterized by the existence of figure 1. typical boiling curve [1]. discrete liquid drops entrained in a continuous vapour flow; normally encountered at void fractions above 80 %. the transition between these two cases is the slug flow film boiling. the existence of film boiling depends on flow conditions and surface temperature. the main parameters influenced heat transfer in post dryout region are pressure, equilibrium quality and mass flux. severe non-equilibrium conditions between liquid and vapour phase can occur and must be taken into account in calculation during low coolant mass flux. it is important to point out that post dryout transition behaviour is reliant on independent boundary condition. if heat flux is the independent boundary condition, the transition boiling regime does not occur. in contrary, when wall temperature is an independent variable, each post dryout flow regime can succeed. as a consequence, the nucleate boiling regime is immediately followed by film boiling regime in the nuclear reactor core [1]. 8 http://dx.doi.org/10.14311/ap.2016.4.0008 http://ojs.cvut.cz/ojs/index.php/app vol. 4/2016 post critical heat transfer and fuel cladding oxidation 2. basic equations first, the main objective of the project was to develop a program for the thermo-hydraulic calculation of nuclear reactor core or experimental electrically heated tubes and rod bundles cooled by light water. the temperature profile of the fuel cladding and fuel with the influence of oxide layer was also included. eventually, the program was used for a comparative assessment of different heat transfer coefficient correlations in the post dryout region and also for evaluation of fuel cladding oxidation correlations. as a conservative approach for the solution of this calculation is used a method of isolated channel model. the calculation can be conducted for stationary states as well as for transients (pseudo-stationary model). it is possible to input nonuniform power distribution and also local resistance (spacer and mixing grids). the program considers homogeneous equilibrium model (hem) which respects thermodynamic imbalance between phases. the calculation can be run at constant pressure or with respect to pressure drops. there are included pressure drops caused by friction, spacer and mixing grids, change of elevation and change of coolant density. pressure drops related to friction between coolant and wall are determined by fiolenko correlation [2]. the wall temperature correction is set by protopov correlation [2] and two phase region correction is calculated by armand’s model [2]. the calculation is based on the solution of equation h = ∫ l 0 q(z) dz. (1) the thermo-physical properties of light water are evaluated by a library using iapws-if97 definition. the program also contains equations and correlations for real quality, slip ratio, void fraction etc. during the development, the code was aimed to a simplicity of input data entering and an ability of easy addition of new models and correlations. 3. heat transfer the fuel cladding surface temperature is very important parameter in operation and safety of nuclear reactors. that is influenced by a local boiling regime and by a suitable correlation for heat transfer from wall to the coolant. in the subcooled liquid region the calculation of nusselt number is conducted by dittus-boelter equation [2] nu = 0.023re0.8pr 0.4. (2) whereas nucleate boiling occurs when wall temperature exceeds coolant saturation temperature by jenslottes difference [2] ∆tjl = 0.791 exp ( − p 6.201 × 106 ) q0.25. (3) the beginning of bulk boiling is defined by saturation temperature of coolant and then thom correlation is used [2] tw = tsat + 22.52 √ q exp ( p 8.6875 ). (4) the wall temperature in the transitional region is linearly approximated. the post critical heat flux heat transfer correlation is used immediately after dnbr decreases below one. the user can choose pg-t [3] or bezrukov [2] critical heat flux correlation. eventually, if equilibrium quality is greater or equal one, convection to superheated steam region is set – sieder-tate correlation nu = 0.021re0.8pr 0.4 ( pr prw )0.25 . (5) 3.1. post critical heat flux heat transfer due to a large number of post critical heat flux heat transfer correlations, it was necessary to carefully select suitable one for use in pressurized light water reactors. the main criterion for selection was validity range of correlation. finally, it was chosen seven correlations for the post critical heat flux heat transfer. a list of chosen post chf heat transfer correlations and their validity ranges are shown in table 1. it is obvious that all correlations are valid in wide ranges and covered most of operational and emergency states of pwr. most of them consider thermo-dynamic equilibrium between the phases, i.e. it assumes that the temperature of vapour phase is equal to the temperature of saturation. in general, these correlations have similar form to dittus-boelter correlation. correlations which respects thermo-dynamic imbalance between the phases, consider superheated steam which coexists with liquid drops. the difference between them is the heat flux calculation using wall temperature and superheated steam temperature. the typical correlation which consider thermo-dynamic imbalance is groeneveld-delorme correlation. correlation p g x [mpa] [kg m−2 s−1] [–] bishop 4.08–21.9 700–3140 0.07–1 groeneveld 5.7 0,07–21,5 130–4000 −0,12–3,09 groeneveld 5.9 3,4–21,5 700–5300 0,1–0,9 miropolskiy 3,9–21,6 800–4550 0,06–1 groe.-delorme 0,7–21,5 130–5200 −0,12–3,09 condie-beng. 0,42–21,5 16,5–5234 −0,2–1,73 pdo tables 0,1–20,0 0–7000 −0,2–2,0 table 1. validity range of used post chf correlations. the first correlation considering thermo-dynamic equilibrium is bishop correlation (6), [1]. the calculation of heat transfer coefficient is applied to the temperature of vapour blanket on the heated surface which can be obtained as the average of wall and 9 vojtěch caha, jakub krejčí acta polytechnica ctu proceedings saturation temperature. this correlation is used for example in vipre or fraptran codes. nu = 0.0193re0.8vf pr 1.23 vf ( %v %l )0.068( x + %v %l (1 −x) )−0.68 (6) the next usually used equation is groeneveld correlation (7), [4]. the user can find to versions (5.7 and 5.9) in the program. the versions differ in coefficients and validity ranges. this correlation is used in cobra-flx and fraptran codes. the temperature of vapour and liquid are related to the saturation temperature. nu = c1 ( gxdh αλv )aprbw( 1 − 0.1 (( %l %v − 1 ) (1 −x) )0.4)−c (7) the program also contains miropolskiy correlation (8), [5]. it is also correlation which respects thermodynamic equilibrium between the phases. but it is valid only for 0.23 ≤ q ≤ 1.16 mw/m2 and 8 ≤ dh ≤ 24 mm. nu = 0.023 ( gdh λv )0.8pr 0.8w (x + %v%l (1 −x))0.8( 1 − 0.1 (( %l %v − 1 ) (1 −x) )0.4)−1 (8) groeneveld-delorme correlation (9), [6] is a typical example of correlation which considers thermodynamic imbalance. it is used for example in fraptran code. nu = 0.008348pr 0.6112vf( gdh µvf ( xa + %v%l (1 −xa) ))−0.8774 (9) the last correlation which was used in the program is condie-bengston correlation (10), [6]. it is an empiric correlation that respects thermo-dynamic equilibrium between the phases. α′ = 5.345 × 10−5 pr 2.2598w re (0.6249+0.2043 ln(x+1)) v (103λv)−0.4593d0.8095h (x + 1)2.0514 . (10) moreover, groeneveld’s look-up post dryout tables [7] are included in the program. its advantage is in the wide range of validity due to the large experimental database which is used and simply searching of heat transfer according to pressure, mass flux equilibrium quality and wall and saturation temperature difference. besides, basic table model includes corrections for hydraulic diameter, cold wall, narrow gap or local resistances. these tables are used for example in assert-pv code. 4. temperature profile and cladding oxidation the module solving fuel temperature profile and cladding oxidation is also included. first, the coolant and wall temperature calculation is performed. an increase of cladding oxide layer and its influence on cladding temperature calculation is followed. then, a conduction in the cladding is solved. heat transfer in fuel-cladding gap calculation is the next step. finally, the temperature profile in fuel is found. all used models consider a cylindrical geometry. the temperature increase caused by lower thermal conductivity of zro2 compared to cladding depends on oxide thickness and its thermal conductivity (function of temperature) ∆tox = qdox λox . (11) the growth of the oxide layer greatly depends on the temperature, cladding material and chemical composition of coolant. the current version of the program contains correlation for low temperature oxidation [8] from matpro library, but is prepared for addition another formulae ∆dox = 0.175 exp ( − 14080 t ) t. (12) heat conduction in the cladding is provided with standard equation of fourier’s law in cylindrical geometry. the calculation of the heat transfer coefficient in fuel-cladding gap is the most complicated part in the temperature profile determination. state of gap depends on many parameters. the program solves it as a sum of three parts – heat conduction in gas, in fuel-cladding contact and thermal radiation. all parts contain several models and many constants. a detailed description can be found in [9]. the fourier’s law is used also for calculation of temperature profile in the fuel [9]. the heat source is considered radially uniform. thermal conductivity in fuel respects dependency on temperature, porosity, burnup, cracking, composition etc. 5. post chf experiments there were carried out a whole range of experiments in post dryout region in tubes as well as in rod bundles which most of them are unavailable due to their commercial utilization. available experiments were carried out at different input parameters (pressure, mass flux, power) and geometries. a comprehensive research was done in [1]. a comparative analysis of post dryout heat transfer coefficient correlations described in section 3.1 were performed. owing to wide range of experimental parameters were chosen experiments measured at the royal institute of technology in stockholm [3]. in these experiments which were carried out over 500 were gained more than 15000 points in post chf region. the experimental equipment was connected to light water loop and was in the shape of a vertical round tube that was electrically heated. there were 10 vol. 4/2016 post critical heat transfer and fuel cladding oxidation also placed 49 thermocouples on the outer wall at different elevation. the axial power distribution was uniform. ranges of experimental parameters are given in table 2. heated length [m] 7.0 inner diameter [m] 0.0100, 0.0149, 0.0247 inlet coolant subc. [◦c] 5 ± 1, 10 ± 3 pressure [mpa] 3, 5, 7, 10, 12, 14, 16, 18, 20 mass flux [kg m−2 s−1] 500, 1000, 1500, 2000, 2500, 3000 heat flux [mw m−2] 0.10–1.25 quality (pdo) [-] 0.03–1.60 table 2. range of experimental parameters and geometry. 6. post chf experiments results the comparison of all described post chf heat transfer coefficient correlation was performed on the group of 120 experiments which were chosen carefully to satisfy validity ranges of all correlations (table 1). the calculation found wall temperatures which were compared with experimentally obtained values in the post chf region. an example of one experiment results is given in figure 2. the figure shows a dependence of wall temperature on the axial length. immediately after chf conditions are reached the rapid increase in wall temperature is observed. then, the wall temperatures differ due to used correlation. figure 2. dependence of calculated and experimental wall temperature on axial length. another capability of developed program is shown in figure 3. the graphical module is able to draw any important calculated parameter in time (transient calculations) and also temperature profiles of coolant, cladding and fuel in each time step. the best results were obtained with groeneveld’s look-up pdo tables. statistical evaluation is given in table 3 where ∆i = tw,calc,i −tw,exp,i tw,exp,i , (13) figure 3. coolant, cladding and fuel temperature profile. ∆̄ = 1 n n∑ i=1 ∆i, (14) σ̄ = √√√√ 1 n − 1 n∑ i=1 (∆i − ∆̄)2. (15) correlation ∆̄ [%] σ̄ [%] bishop 2.22 8.69 groeneveld 5.7 0.19 8.3 groeneveld 5.9 −4.93 7.37 miropolskiy 12.58 22.37 groe.-delorme 92.01 12.08 condie-beng. 13.73 13.69 pdo tables 2.82 5.67 table 3. statistical evaluation of post chf correlations. figure 4 shows a comparison of experimental and calculated wall temperatures achieved with groeneveld’s look-up pdo tables. figure 4. groeneveld’s look-up pdo tables results. 11 vojtěch caha, jakub krejčí acta polytechnica ctu proceedings 7. conclusion a thermo-hydraulic program based on isolated channel model was developed. it includes the wide range of heat transfer correlations for different one phase and two phase flow regimes. great attention was paid to the post chf boiling regime. the program also includes the module for evaluation of temperature profile in fuel and cladding which contains influence of oxide layer. the program is prepared for the addition of new oxidation model correlations which are being developed for different cladding tubes materials and temperature ranges. all calculated results can be shown in 3d graphs created by post-processing module using python mayavi library. a comparative analysis of different heat transfer correlations in the post chf region was done. the analysis included a comparison of seven post chf heat transfer correlations with experimental data which were obtained on electrically heated tube. the best results were achieved with groeneveld’s look-up pdo tables. the continue of this work should aim to verify this results on different experimental data set with axially nonuniform power distribution, on transient experiment with the return from post chf region and on rod bundle experiment. list of symbols g mass flux [kg/m2s] h enthalpy [k/kg] l length [m] q heat flux [w/m2] q linear power [w/m] nu nusselt number [–] p pressure [pa] pr prandtl number [–] re reynolds number [–] t time [s] t temperature [k] x quality [–] α void fraction [–] λ heat transfer coefficient [w/m2k] µ dynamic viscosity [pa s] % density [kg/m3] acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs14/156/ohk4/2t/14. references [1] thermohydraulic relationships for advanced water cooled reactors. tech. rep. iaea-tecdoc-1203, international atomic energy agency, 2001. http://www.iaea.org/inis/collection/ nclcollectionstore/_public/32/024/32024154.pdf. [2] n. e. todreas, m. kazimi. nuclear systems volume i: thermal hydraulic fundamentals. crc press, 1989. [3] k. m. becker, c. h. ling, s. hedberg, g. strand. an experimental investigation of post dryout heat transfer. tech. rep. kth-nel-33, royal institute of technology, stockholm, sweden, 1983. http://www.iaea.org/inis/collection/ nclcollectionstore/_public/15/049/15049472.pdf. [4] d. c. groeneveld. post-dryout heat transfer at reactor operating conditions. in national topical meeting on water reactor safety. 1973. http://www.iaea.org/inis/collection/ nclcollectionstore/_public/04/089/4089010.pdf. [5] n. hammouda. subcooled film boiling in non-aqueous fluids. ph.d. thesis, university of ottawa, 1996. http://hdl.handle.net/10393/9949. [6] s. k. moon, s. y. chun, s. cho, et al. an experimental study on post-chf heat transfer for low flow of water in a 3x3 rod bundle. nuclear engineering and technology 37(5):457–468, 2005. [7] d. c. groeneveld. look-up table for fully developed film-boiling heat-transfer coefficients for light water. university of ottawa, 2001. [8] j. krejčí, v. vrtílková, d. gajdoš, d. rada. proposal of new oxidation kinetics for sponge base e110 cladding tubes material. in topfuel2015 – conference proceedings, pp. 466–473. 2015. http://www.euronuclear.org/ events/topfuel/topfuel2015/transactions/ topfuel2015-transactions-poster.pdf. [9] j. krejčí, v. caha. ube-postchf. tech. rep., čvut, praha, 2015. 12 http://www.iaea.org/inis/collection/nclcollectionstore/_public/32/024/32024154.pdf http://www.iaea.org/inis/collection/nclcollectionstore/_public/32/024/32024154.pdf http://www.iaea.org/inis/collection/nclcollectionstore/_public/15/049/15049472.pdf http://www.iaea.org/inis/collection/nclcollectionstore/_public/15/049/15049472.pdf http://www.iaea.org/inis/collection/nclcollectionstore/_public/04/089/4089010.pdf http://www.iaea.org/inis/collection/nclcollectionstore/_public/04/089/4089010.pdf http://hdl.handle.net/10393/9949 http://www.euronuclear.org/events/topfuel/topfuel2015/transactions/topfuel2015-transactions-poster.pdf http://www.euronuclear.org/events/topfuel/topfuel2015/transactions/topfuel2015-transactions-poster.pdf http://www.euronuclear.org/events/topfuel/topfuel2015/transactions/topfuel2015-transactions-poster.pdf acta polytechnica ctu proceedings 4:8–12, 2016 1 introduction 2 basic equations 3 heat transfer 3.1 post critical heat flux heat transfer 4 temperature profile and cladding oxidation 5 post chf experiments 6 post chf experiments results 7 conclusion list of symbols acknowledgements references 20 acta polytechnica ctu proceedings 1(1): 20–26, 2014 20 doi: 10.14311/app.2014.01.0020 gravitational waves and dark energy peter l. biermann1,2,3,4,5, benjamin c. harms1 1department of physics and astronomy, the university of alabama, box 870324, tuscaloosa, al 35487-0324, usa 2mpi for radioastronomy, bonn, germany 3karlsruhe institute of technology (kit) institut für kernphysik, germany 4department of physics, university of alabama at huntsville, al, usa 5department of physics & astronomy, university of bonn, germany corresponding author: bharms@bama.ua.edu abstract the idea that dark energy is gravitational waves may explain its strength and its time-evolution. a possible concept is that dark energy is the ensemble of coherent bursts (solitons) of gravitational waves originally produced when the first generation of super-massive black holes was formed. these solitons get their initial energy as well as keep up their energy density throughout the evolution of the universe by stimulating emission from a background, a process which we model by working out this energy transfer in a boltzmann equation approach. new planck data suggest that dark energy has increased in strength over cosmic time, supporting the concept here. the transit of these gravitational wave solitons may be detectable. key tests include pulsar timing, clock jitter and the radio background. keywords: cosmology dark energy black holes gravitational waves. 1 introduction dark energy was originally detected as accelerated expansion seen in the distance scale for supernovae of type ia (schmidt et al., 1998, riess et al. 1999, perlmutter et al. 1999; for a review see frieman et al. 2008). many suggestions have been made about what dark energy is, what its strength is, what its time evolution is, and what possible further observational results are. the idea that dark energy is gravitational waves may explain its strength and its time-evolution. one possible concept is that dark energy is the ensemble of coherent bursts (solitons) of gravitational waves originally produced when the first generation of supermassive back holes was formed (caramete & biermann 2010); the energy density of such solitons would suffice within the uncertainties. these solitons get their initial energy as well as keep up their energy density throughout the evolution of the universe by stimulating emission from a background (biermann & harms 2013). our model of the background metric resembles the randall-sundrum ideas (1999a, b) but is timedependent, and describes the energy flow from the background (strong-gravity) brane to our world (weakgravity) brane. planck data suggest that dark energy has increased in strength over cosmic time (planck 2013 xvi), as predicted by our model. gravitational waves were far below today’s dark energy at the epoch of early nucleosynthesis and of the formation of the microwave background ripples (as summarized in ligo+virgo-coll. 2009), both much earlier than the likely formation epoch of the first generation of super-massive black holes. our model is also consistent with early star formation (biermann et al. 2014), as we argue below. the transit of the gravitational wave solitons postulated here may be detectable. we discuss the predictions briefly below and elsewhere. we focus on the boltzmann equation approach, working out the energy transfer from the strong gravity background in stimulated emission. 1.1 gravitational solitons from black hole mergers inspired by bekenstein’s (1973) considerations we posit: when the first generation of super-massive black holes was formed, each produced a coherent burst of solitonlike gravitational waves which combine to give a total energy of order ∼ 1 2 nbh,0 mbh c 2 (1 + z?) 3 . (1) in the following we also call this an ensemble of soliton waves, or shell fronts. nbh,0 is the original comoving density of super-massive black holes. today supermassive black holes have a density of 10−1.7±0.4 mpc−3 20 http://dx.doi.org/10.14311/app.2014.01.0020 gravitational waves and dark energy above mbh = 3·106 m� (caramete & biermann 2010); assuming that they grow by merging, and allowing for statistical and systematic errors, an original comoving density of nbh,0 = 1 mpc −3 seems possible. this comoving density is the density black holes had at the beginning, so transposed to today without change in their numbers per comoving volume. the data suggest that there was a generation of first super-massive black holes with a mass between mbh ∼ 106 m� and mbh ∼ 107 m�. the original black hole mass may be ∼ 3 · 106 m� considering (i) the black hole mass function (greene et al. 2006, caramete & biermann 2010), (ii) the instability of massive stars (appenzeller & fricke 1972a, b) in an agglomeration picture (spitzer 1969, sanders 1970), and (iii) the observed black hole in our galactic center (e.g. eckart et al. 2005). the redshift of creation z? may be large, as formation of massive stars may begin at redshift 80 (biermann & kusenko 2006). redshifts z? from about 30 to 50 allow a quantitative interpretation of the data of dark energy. at the original density of black holes adopted here redshift 50 is consistent with the mass of mbh ∼ 3 ·106 m�, and redshift 30 would imply mbh ∼ 107 m�, in either case to make the estimate consistent with dark energy today. what is the motivation for considering gravitational waves? bekenstein (1973) wrote about the entropy of the universe: “... we must regard black hole entropy as a genuine contribution to the entropy content of the universe”. however, entropy is also information, and information must have a carrier. a natural suggestion is that this carrier is gravitational waves, with an energy commensurate with the black hole scale. this suggestion is consistent with the fact that cosmological black holes are not in thermodynamic equilibrium, and therefore the entropy associated with such black holes should be described by statistical mechanics as advocated in harms and leblanc (1992, 1993). this speculation immediately gives s kb = ngw,0 = 4π ( mbh mpl )2 , for zero spin , (2) where ngw,0 is the number of gravitons at the formation of the black hole. for mbh = 3 · 106 m� this is ngw,0 ' 1090. mbh is the original mass of the black hole, mpl is the planck mass, c is the speed of light, and gn is newton’s constant of gravity. egw is the average graviton energy given by egw = 1 8π h̄c3 gnmbh = c2 8π m2pl mbh . (3) this gives a graviton energy of egw ' 10−30 erg for this black hole mass. the entire energy content then is ngw,0 egw = 1 2 mbh c 2 . (4) we picture this as a coherent burst of gravitational waves, or a soliton wave, ejected at formation of the black hole. it is clear from the considerations above that we are not using the weak-field approximation. multiplying with the original density of super-massive black holes reproduces our estimate above. 1.2 five-dimensional background model in our model for the background, which has some similarity to the randall-sundrum (1999a, b) ideas, we identify a possible local metric to describe a 5d world with a 4d strong gravity brane and our 4d world weakgravity brane. ds2 = −e(u/l) m t/ψ c2 dt2 + e( u l ) p t/β du2 + e(1−b( u l ) n ) 2 t/αth e−( u l ) k (1− tφ ) dxi dx i , (5) where i = 1, 2, 3, τh is the hubble time, l = lpl is the planck length, τpl = lpl/c is the planck time, u is the coordinate in the fifth dimension, and the remaining, non-coordinate quantities are arbitrary parameters. although the five-dimensional covariant divergence of the energy-momentum tensor does not vanish everywhere, it vanishes on the weak-gravity (weak) brane (u = 0), and it is approximately zero on the strong-gravity (strong) brane (u = l) for very small dimensionless ratios β/τh and ψ/τh with ψ < 0 and β > 0, with the conditions a) τh(ψ + β)/(ψβ) >> 1, b) 2(b−1)/α > τh/φ−1, and c) α2 = 3. g00 in the metric above then defines the confining potential for the strong brane. gravitons from our brane stimulate transitions between bound states on the strong brane, resulting in the emission of gravitons onto our brane. this metric describes a weak brane for our world which is expanding with time, and a strong brane which is contracting with time, albeit very slowly for the latter brane. the 5d-cosmological constant measured on the weak brane is ωweak = − ( τpl τh )2 . figure 1: the stimulated emission of gravitons in a shell in our model the gravitons on the strong-gravity brane obey a planck-like distribution with planck temperature. we adopt the point of view that the planck scales are limits: nothing can go below planck time 21 peter l. biermann, benjamin c. harms and planck length, and no single particle can go beyond planck energy, in any frame. the strong brane is stable against collapse, since given a planck spectrum for any wavelength λ the free-fall time scale τff is always either equal or longer than the pressure wave time scale τs. figure 2: the strong-gravity (planck brane) and weak-gravity (tev brane) branes τff = τpl ( λ lp l )3/2 ≥ τpl ( λ lp l ) = τs. 2 stimulate emission of energy from the strong-gravity brane in the following we use a particle-wave duality for gravitons at high energy, which thus associates a wavenumber ~k/h̄ and a corresponding length-scale λ to each spatial direction, and we assume localization is possible to about a wavelength. 2.1 rate of energy transfer the distribution function n(k,t) is the distribution of occupied allowed states on our brane for gravitons with momenta k = |~k| and p = |~p| at time t on the shell and satisfies the equation( ∂ ∂t − ṙ(t) r(t) k ∂ ∂k ) n(k,t) = 1 k ∫ d3 k′ (2π)3 2 k′ ∫ cd3 p (2π)3 2 e(p) ∫ cd3 p′ (2π)3 2 e(p′) ∫ d3 k′′ (2π)3 2 k′′ γ2 |m|2 (2 π)4δ4(k + q−k′ −k′′ −q′) (2 π)6 δ3(~k′ − ~k′′) δ3(~k − ~k′) [gb(p ′, t)(1 + n(k,t)) ((n(k′, t) + 1)(n(k′′, t) + 1) − 1) −n(k,t)gb(p,t)(1 + n(k′, t))(1 + n(k′′, t))] . (6) where k = (k,~k), γ = k3ref,1, and q = (e,~p). kref,1 is a reference momentum to be determined below. the δ-functions, δ3(~k′ − ~k′′) and δ3(~k − ~k′), have been inserted to impose coherence of the outgoing gravitons. |m|2 is the matrix element squared for the quadrupole emission of a graviton of 4-momentum k′′, and has the dimensions of (momentum)−3 (time)−1. we will assume that this matrix element squared is proportional to k5. gb(p,t) is the occupation number distribution of the background particle sea. r(t) = (1 + z?)/(1 + z) is the scale factor for an expanding universe. the following analysis is done in the observer frame. the boltzmann equation for n(k,t) to lowest order in the expansion of the 4-dimensional δ-function is( ∂ ∂t − ṙ(t) r(t) k ∂ ∂k ) n(k,t) ' +κ k n(k,t) (n(k,t) + 1) , (7) where the factor κ is given, after integration over all the δ-functions, by κ = k2ref,2 h(z) 24πkbh |m|2 ln{ kbh+ k } . (8) in the equation above kbh+ is the maximum momentum at which stimulated emission of gravitons occurs, just above the momentum of the peak of n(k,t). since the log-term in eq.8 varies very slowly over the range of k of interest we approximate this term with a constant and set β = {ln(kbh+/k)}/(24 π). |m|2 is related to |m|2 by extracting the factors (k/kref,1)3/k3ref,1, h(z). a threshold function of 2 gbn/(gb + n) arises from the normalized interaction between the gravitons on the background brane and our brane; this function connects to the strong brane only if n > gb, which is the condition for stimulated emission. these choices do not introduce new constraints or additional assumptions. next we redefine (k/kref,2) 2 κ = κ to extract the k-dependence from |m|2. making the change of variables k = k̃/r(t), eq.7 can be written as ∂n(k̃, t) ∂t ' h(z) k̃ β kbh r(t) n(k̃, t) (n(k̃, t) + 1) . (9) in terms of the frequency of the wave at emission this equation is ∂n(ν0, t) ∂t ' h(z) hν0 β kbh r(t) c n(ν0, t)(n(ν0, t) + 1) . (10) introducing the dimensionless variables x = hν0 kb tg0 , and y = ∫ t 0 kbtg0 h(z)β kbh cr(t) dt′ , (11) where kbhc = kbtg0 = mplc 2 mpl 8 π mbh . eq.(10) becomes ∂n ∂y ' +xn(x,y) (n(x,y) + 1) . (12) the solution of this equation is n(x,y) = 1 ex(a−y)+b − 1 , (13) 22 gravitational waves and dark energy where a and b are constants, to be determined later. this distribution (eq.13) is planck-like with a timedependent normalized temperature 1/a. the rate at which energy is created can be calculated from the expression for n in eq.10. the rate at which energy is created can be calculated from the expression for n in eq.(10). the rate of energy creation per existing graviton (of the total number ngw,0 r(t) 4) is d < e > dt = βh(z) r(t) × (14)∫ x3 hν0 n(ν0, t) (n(ν0, t) + 1) dx. the total rate of energy creation is d < et > dt = ngw,0 r(t) 3 kbh ch(z) β a , (15) where a = ∫ x4 n(x,t) (n(x,t) + 1) dx (16) inserting all these constants into the integral for y demonstrates that y approaches a constant for the redshift z? being large, and integrating down to today or even into the future, when y approaches a constant of β << 1. a is equivalent to an inverse temperature, and should be of order unity. without loss of generality we can set b = 0. this integral strongly depends on the exact value of a − y, and is of order 30 for a − y ' 1, and b approaching zero. the matrix element |m| does not evolve with time, and scales as momentum |m| = �m ( k mplc ) . (17) above we have used �m = 1; we now generalize and allow �m to be different from unity. writing |m| in this way suggests that the interaction between the gravitons on our brane and the gravitons on the background brane comes down to a fundamental coupling constant. this behavior is consistent with the idea [22], that the gravitational coupling strongly increases with energy to approach the other three coupling constants at near planck energies. this allows the expression for d<et> dt to be consistent with the observed energy density under the condition that aβ �m = 3. for the constant a = 1 above, β of order 0.1, and �m = 1, the quantity {aβ �m} is in fact close to 3. however, if we were to require that the k-range be very large, then β would be larger, and �m would be required to be smaller than unity accordingly. inserting this parameter dependence into eq.(15) then leads back, to within the approximation that aβ �m = 3, to the result we were seeking, 3 2 mbh c 2 h(z) ( 1 + z? 1 + z )3 . (18) after integrating we obtain with this redshift dependence a constant dark energy density as in eq.( 1) by multiplying by the redshift evolution of black holes nbh,0 (1 + z) 3. the factor of 4 multiplying ρde in eq.22 below derives from the sum of dark energy density and pressure, and corresponds to 3 (ρde +pde/c 2). therefore the rate of change of dark energy with time is 3 (ρde + pde/c 2) h(t) and today 3ρdeh(z = 0) = 3 2 mbh c 2 h(z = 0) (1 + z?) 3 . (19) this justifies the ρde h(t) term in eqs.22 and 25. this shows that dark energy remains at the level of eq.(1) throughout the evolution of the universe, in the approximation that most early super-massive black holes were formed over a short span of time. 2.2 equation of state we define ρde(t,u) as the dark energy density, ρ(t,u) as the total energy density, p(t,u) as the total pressure, and we use the equation of state pde(t,u) = ρde(t,u) c 2/3 . the dark energy density ρ(t,u) is scaled to the value of the dark energy density observed on our brane today ( redshift z = 0 ). the friedmann-robertson-walker form of the einstein equations on the weak-gravity brane must be (h(t)) 2 = ( ˙r(t) r(t) )2 = 8 π gn 3 ρ(t, 0) (20) and r̈(t) r(t) = − 4 π gn 3 ( ρ(t, 0) + 3 p(t, 0) c2 ) + 16π gn 3 ρde(t, 0) + 4π gn sinj 3 h(t) , (21) where the rate of change of the energy density used is ρ(t, 0) = − 3 ( ρ(t, 0) + p(t, 0) c2 ) h(t) + 4 ρde(t, 0) h(t) + sinj . (22) we emphasize that in the second equation, eq.21, the term 16 π gn ρde(t, 0)/3 corresponds to the continuous energy transfer by stimulated emission on the basis of existing dark energy; the additional term (4 π gn sinj)/(3 h(t)) describes new formation of dark energy. this allows a different equation of state for exactly the same cosmological observations, as now for pde = ρdec 2/3 the modified equation 21 becomes identical to the canonical version of this equation for 23 peter l. biermann, benjamin c. harms pde = −ρdec2. the corresponding set of equations for the strong-brane are hsb(t) 2 = 8 π gn 3 (ρde(t, lsb) − λsb) (23) where λsb is the cosmological constant on the strong brane at the beginning of the epoch of black hole formation and r̈sb(t) rsb(t) = − 16 π gn 3 ρde(t, 0) h(t) hsb(t) (24) + 8 π gn 3 ( 3 ρde(t, lsb) − λsb) − 4 π gn sinj 3 hsb(t) , the rate of change of the energy density on the strongbrane is correspondingly ρ̇de(t, lsb) = 3 ( ρde(t, lsb) + 1 c2 pde(t, lsb) ) hsb(t) −4 ρde(t, 0)h(t) −sinj . (25) eq.24 is derived from eq.23 using the conservation of energy-momentum equations; eqs. 22 and 25 insure that energy is conserved between the two branes, the 4-dimensional boundaries of the 5-dimensional universe described by our model. for epochs before the energy transfer started ρde(t, 0) = 0, and sinj = 0, and we can set ρde(t, lsb) = λsb for t < t(z = z?). 3 comparison to experimental limits our model can be tested by several different types of experiment. we discuss two of these below. 3.1 pulsar timing experiments figure 3: gravitational wave background limit from pulsar timing (dashed line) , and our inferred gravitational wave background from stimulated emission of gravitational waves from the background planck sea constituting dark energy (straight line). the ordinate is the fraction of closure density ω per log bin of frequency, and the abscissa is the frequency of the gravitational waves f. the gravitational waves in our model arise from the production of black holes in the early universe at a redshift of z ' 50. the observed constant dark energy density is maintained by the continuous production of gravitational waves by black-hole interactions with the planck sea background. for an idealized model in which all black holes were created at the same time, and with the same mass, the gravitational wave background peaks near fgw,max ' 10−4.5 hz m−1bh,6.5 (1 + 50)/(1 + z?). a soliton comes past a given point in space-time on the order of every 20 seconds. wave forms are created by uncorrelated solitons passing by a given point. ultra-precise timing experiments which test the steadiness of timing over time scales of order a few seconds to a few minutes would show these variations in the energy density if the precision is high enough. the precision to detect such a signal has to correspond to seconds in the expansion rate of the universe, which requires the precision to be of order 10−17.5, or a few 10−18. this precision is expected to be reached in the next generation of clocks [16]. t ρ 1 figure 4: the sequential passing of solitons (sharp peaks) can be approximated by a sinusoidal wave form. 4 conclusions if the validity of our model is proven by experimental tests such as the pulsar timing experiments (fig.3), the detection of time jitter (fig.4), or the detection by an observatory such as ligo or virgo of the passage of a shell front by a dedicated type of data analysis, the implications for cosmology are great. the basis of our model is that the source of dark energy is the creation of gravitational waves by the interaction of surfaces at critical density, e.g. the planck surfaces surrounding black holes, with a strong-gravity brane located a few planck lengths from our weak-gravity brane. our model is consistent with the big-bang theory after the first planck time. however, in our model the universe starts from a lemaitre-like ‘atom’ or ‘seed’ [5]. our model also has implications for quantum gravity theory. experimental evidence for the validity of 24 gravitational waves and dark energy our model would imply the existence of a strong-gravity brane and extra dimensions as well as a smallest distance and impenetrable planck surfaces rather than horizons. although our model describes several observed cosmological phenomena, it is largely heuristic. we are currently working on an exact solution of the fivedimensional space-time metric and a more formal mathematical description of the stimulated emission amplitude for the creation of the solitons. acknowledgement discussions with lou clavelli greatly contributed to the development of the paper; discussions with laurenţiu caramete (bucharest, romania), roberto casadio (bologna, italy), marco cavaglia (oxford, ms), laszlo gergely (szeged, hungary), shaoqi hou (tuscaloosa, al), pankaj joshi (mumbai, india), gopal-krishna (pune, india), octavian micu (bucharest, romania), piero nicolini (frankfurt, germany), norma sanchez (paris, france), joe silk (oxford, united kingdom), allen stern (tuscaloosa, al), dijan stoikovic (buffalo, ny), and hector de vega (paris, france) are gratefully acknowledged. helpful comments on earlier versions of the manuscript were received from r. casadio, m. cavaglia, l. gergely, a. graham, p. joshi, o. micu, p. nicolini, and d. stoikovic. this research was supported in part by the doe under grant de-fg02-10er41714. references [1] appenzeller, i., & fricke, k., a&a 18, 10 (1972a); appenzeller, i., & fricke, k., a&a 21, 285 (1972b) [2] bekenstein, j.d., phys. rev. d 7, 2333 (1973) doi:10.1103/physrevd.7.2333 [3] biermann, l., schlüter, a., phys. rev. 82, 863 (1951) doi:10.1103/physrev.82.863 [4] biermann, p.l., kusenko, a., phys. rev. lett 96, 091301 (2006) doi:10.1103/physrevlett.96.091301 [5] biermann, p.l., & harms, b.c., eprint arxiv:1205.4016 (2012); biermann, p.l., & harms, b.c., in proc. at the 13th marcel grossmann meeting, july 2012, eprint arxiv:1302.0040 (2013) [6] biermann p.l., et al., mon. not. roy. astron. soc. 441, 1147 (2014) doi:10.1093/mnras/stu541 [7] caramete, l.i., & biermann, p.l., a&a 521, id.a55 (2010) [8] eckart, a., et al., in the evolution of starbursts: the 331st wilhelm and else heraeus seminar. aip conf. proc. 783, 17 (2005) [9] frieman, j.a., turner, m.s., & huterer, d., ara&a 46, 385 (2008) doi:10.1146/annurev.astro.46.060407.145243 [10] greene, j.e., barth, a.j., & ho, l.c., new astro. rev. 50, 739 (2006) doi:10.1016/j.newar.2006.06.080 [11] harms, b.c., & leblanc, y., phys. rev. d 46, 2334 (1992); harms, b.c., & leblanc, y., phys. rev. d47, 2438 (1993) [12] ligo and virgo collaborations, nature 460, 990 (2009) doi:10.1038/nature08278 [13] munyaneza, f., & biermann, p.l.,a&a 436, 805 (2005); munyaneza, f., & biermann, p.l., astron. & astroph. letters 458, l9 (2006) [14] perlmutter, s., et al., apj 517, 565 (1999) [15] planck collaboration; ade, p.a.r., et al., eprint arxiv:1303.5078 (2013) [16] predehl, k., et al., science 336, 441 (2012) [17] randall, l., sundrum, r., phys. rev. lett 83, 3370 (1999a); randall, l., sundrum, r., phys. rev. lett 83, 4690 (1999b) [18] riess, a.g., et al. aj 116, 1009 (1998) [19] sanders, r.h., apj 162, 791 (1970) [20] schmidt, b.p., et al., apj 507, 46 (1998) [21] spitzer, l., jr., apjl 158, l139 (1969) doi:10.1086/180451 [22] dimopoulos, s, raby, s.a., wilczek, f. phys. today oct., p. 25 (1991) discussion moshe elitzur: are there predictions for the effect of your model on the imprint of fluctuations on cmb and baryon acoustic oscillations? benjamin harms : dark energy in our mode is due to the merging of black holes at a ' 50, so well after the formation of the cmb. there may be some effect on the propagation of the cmb photons, but we have not yet worked out the exact nature of this effect. 25 http://dx.doi.org/10.1103/physrevd.7.2333 http://dx.doi.org/10.1103/physrev.82.863 http://dx.doi.org/10.1103/physrevlett.96.091301 http://dx.doi.org/10.1093/mnras/stu541 http://dx.doi.org/10.1146/annurev.astro.46.060407.145243 http://dx.doi.org/10.1016/j.newar.2006.06.080 http://dx.doi.org/10.1038/nature08278 http://dx.doi.org/10.1086/180451 peter l. biermann, benjamin c. harms jim beall: can you comment on the effect these ’seed’ black holes have on galaxy formation? benjamin harms: the main difference between our model and the ’big bang’ theory is that our model allows for large assemblies of stars which have never merged and do not have an agn, which have apparently been observed. 26 introduction gravitational solitons from black hole mergers five-dimensional background model stimulate emission of energy from the strong-gravity brane rate of energy transfer equation of state comparison to experimental limits pulsar timing experiments conclusions acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0107 acta polytechnica ctu proceedings 4:107–112, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of s-co2 power cycles for nuclear energy ladislav veselya, ∗, vaclav dostala, slavomir entlerb a faculty of mechanical engineering, czech technical university in prague, technická 4, 166 07 prague 6, czech republic b institute of plasma physics of the cas, za slovankou 3, 182 00 prague 8, czech republic ∗ corresponding author: ladislav.vesely@fs.cvut.cz abstract. the supercritical carbon dioxide (s-co2) is a possible cooling system for the new generations of nuclear reactors and fusion reactors. the s-co2 power cycles have several advantages over other possible coolants such as water and helium. the advantages are the compression work, which is lower than in the case of helium, near the critical point and the s-co2 is more compact than water and helium. the disadvantage is so called pinch point which occurs in the regenerative heat exchanger. the pinch point can be eliminated by an arrangement of the cycle or using a mixture of co2. this paper describes the s-co2 power cycles for nuclear fission and fusion reactors. keywords: s-co2, power cycle, fusion. 1. introduction the supercritical carbon dioxide (s-co) cycles are recently very prospective power cycles for different applications. these applications are ranging from nuclear through geothermal, solar energy and waste heat recovery systems. these cycles are researched all around the world. the research of the power cycles with co2 as working medium has a long history. the first reference is dated back to 1948, when sulzer bros. patented a brayton cycle with the partial condensation of co2. the worldwide research of s-co2 power cycles is dating to the second half of the 20th century [1]. researchers began to realize benefits of co2 as working medium in power cycles at that time. among the first researchers who studied benefits of co2 power cycles belongs angelino, feher, verhivker and gokhstein [2]. the research on czech technical university in prague (ctu) is oriented on the analysis of the s-co2 power cycles with a potential for the nuclear reactors, as well as for the fusion power reactors. the design of the s-co2 power cycle is very important and it has effect on the cycle efficiency and net power. the s-co2 cycle is also suitable for utilization of heat from multiple heat sources with different temperature and heat power. this is important in a case of multiple heat sources providing high and low potential heat like fusion power reactors. this paper is focused on comparison of the s-co2 cycles for nuclear energy. benefits of the s-co2 cycle will be described for the nuclear fission and fusion reactors. the design of the s-co2 cycle will be applied for multiple heat source in fusion reactor. 2. advantages and disadvantages of s-co2 cycles the main advantage of the s-co2 cycles is the compression work which is lower than in case of helium [2]. a compressor work reduction is caused due to operation in near the critical point. the critical point of co2 occurs at the temperature 30.98 ◦c and pressure 7.32 mpa. the s-co2 cycle is more compact than water and helium because this cycle operating at high pressure and allows small size of components. another advantage is that the s-co2 cycle achieves high efficiency with low operates temperature. the s-co2 cycles also have several disadvantages. the existence of so called “pinch point” in heat exchangers significantly affecting their design is the most important and well-known disadvantage. the pinch point may be present for any type of medium, but its influence on components is especially high when co2 is employed as a working medium. the pinch point primarily occurs in recuperative heat exchangers with identical working media and mass flow on both the hot and the cold side. the pinch point is caused by the variations of heat capacity of co2 and occurs when the heat capacity of the hot and cold streams (each at a different pressure level) intersect. due to the pinch point, the heat exchangers may have a large size and low efficiency. however, this problem can be removed in several ways. one of them is an addition of the small amount of other substance into the pure co2. the substances for shift of pinch point could be ar, he, co, o2 or n2 [3]. the other ways is a change of the design of cycles and usage of different mass flows in hot and cold side of heat exchangers. 107 http://dx.doi.org/10.14311/ap.2016.4.0107 http://ojs.cvut.cz/ojs/index.php/app l. vesely, v. dostal, s. entler acta polytechnica ctu proceedings 3. description of gas cycles the s-co2 cycle is a gas cycle derived from the ericsson-brayton cycle, which offers many different layouts for solar, geothermal or nuclear power plants and waste heat recovery. each layout tries to approach the carnot cycle and its efficiency. the basic layouts considering the use of s-co2 are [2]: • simple brayton cycle, • re-compression cycle, • pre-compression cycle, • split expansion cycle, • partial cooling cycle, • partial cooling with improved regeneration. figure 1. simple brayton cycle. the simple brayton cycle is shown in figure 1. this cycle has a turbine (t), compressors (c1), recuperative heat exchanger (rh), a cooler (c) and a heater (h). figure 2. re-compression cycle. the re-compresion cycle is shown in figure 2. the difference between the simple brayton cycle and recompression cycle is twice a number of the compressors and recuperative heat exchanger. the re-compression cycle has a turbine (t), two compressors (c1 and c2), two recuperative heat exchangers (ltr and htr), cooler (c) and heater (h). in figure 3, the pre-compression cycle is shown, and figure 4 presents the split expansion cycle. the pre-compression cycle and the split expansion cycle have same components as the re-compression cycle. only two turbines has the split expansion cycle. figure 3. pre-compression cycle. figure 4. split expansion cycle. the partial cooling cycle is shown in figure 5. the partial cooling cycle contains the two coolers (ca and cb) and three compressors. figure 5. partial cooling cycle. 108 vol. 4/2016 comparison of s-co2 power cycles for nuclear energy 4. the gfr re-compressing cycle the s-co2 cycles for nuclear energy are based on the presented cycles of co2. the re-compression cycle applied to the gas-cooled fast reactors (gfr) is analyzed in this study. the re-compression cycle is used as basic concept for gfr. other type of s-co2 cycle has similar results as the re-compression cycle [4]. the cycle layout is arranged according to the figure 2. the ltr exchanger is sensitive to the pinch point due to operation near the critical point. the recompression cycle eliminates the pinch point using different mass flows of the ltr exchanger. the parameters for calculation are shown in table 1. the compressor inlet temperature is 34 ◦c. the turbine inlet temperature is 550 ◦c. the heat source is considered with the minimum inlet temperature into the heater about 600 ◦c. the thermal power is 600 mw [5]. the cycle was optimized for the best parameters with thermal power 600 mw. thermal power 600 mw compressor efficiency 68 % turbine efficiency 79 % recuperator effectiveness 90 % compressor inlet temperature 34 ◦c turbine inlet temperature 550 ◦c table 1. parameter of the gfr s-co2. 5. the s-co2 cycle for demo2 fusion reactor in the case of the fusion reactor, the re-compression cycle is used as the first. the heat sources are arranged behind [6]. however, the layout of the re-compression cycle can be designed differently. the heat sources can be situated to other streams [7]. the different layout of the s-co2 power cycle can be a benefit for the heat transfer and net power of cycles. a demonstration fusion power plant (demo) represents the first fusion power station capable of producing electricity and operating with a closed fuel-cycle. two demo design options are currently investigated, in an attempt to identify a realistic range of possibilities: a near-term demo1 and an advanced design concept demo2. demo1 is the concept based on reliable technology deliverable in the term of 20 years from now, and it is planned to work in the pulse operation mode. demo2 based on advances in the physics basis deliverable on a longer term is expected in the steady-state operation mode [8]. the demo2 power plant based on the steady-state fusion power reactor is analysed in this study. the demo2 fusion reactor has several different heat sources. the main heat sources are the blanket, first wall, and divertor. each of them operates on different temperatures and powers. the preheating s-co2 cycles may be more suitable for demo2. a layout of the preheating cycle is shown in figure 6. the preheating cycle has benefit for the pinch point mitigation as well as the layout of re-compression cycle. figure 6. preheating cycle. the table 2 brings parameters of the analyzed fusion reactor demo2 model [9]. thermal power of the fusion reactor demo2 is 4109 mw. the blanket has the thermal power of 3887 mw. the thermal power of the divertor and the first wall is 222 mw. the blanket and first wall are cooled by helium, the divertor is cooled by water. the high-grade reactor outlet temperature is projected to 500 ◦c. the low-grade reactor outlet temperature is projected to 160 ◦c. the turbine inlet temperature is 475 ◦c. consequently, the analyzed fusion reactor demo2 has two different heat source. the preheating cycle uses the heater h1 for high-grade primary heat (first wall and blanket). second heater h2 is used for lowgrade secondary heat (divertor). thermal power 4109 mw primary (high-grade) heat 3887 mw secondary (low-grade) heat 222 mw compressor efficiency 68 % turbine efficiency 79 % recuperator effectiveness 90 % compressor inlet temperature 34 ◦c turbine inlet temperature 475 ◦c table 2. parameter of s-co2 for fusion reactor demo2. 6. result of gfr reactor and demo2 fusion reactor the thermodynamic calculation was done for precompression cycle and preheating cycle. the calculation was performed using programming language python. the codes of cycles have been written in python. properties of pure co2 and mixtures are embedded into the python. source of gases and mixtures properties is nist reference fluid thermodynamic and transport properties database, version 9.1. [10]. the calculation of the gfr s-co2 cycle was performed according to the parameters included in table 2. the cycle was optimized for the best results. the results of the gfr re-compression cycle are 109 l. vesely, v. dostal, s. entler acta polytechnica ctu proceedings figure 7. t-s diagram of the gfr re-compression cycle. shown in table 3. the figure 7 shows a t-s diagram of the cycle. according to the table 3, the total net power of re-compression cycle is 201 mw. the cycle efficiency is 33.56 % and the mass flow is 2730 kg/s. cycle efficiency 33.56 % turbine power output 324.98 mw compressor no.1 in. power 89.44 mw compressor no.2 in. power 33.84 mw compressor input power 123.28 mw added heat 600.98 mw removed heat 399.28 mw regenerative heat 1249.15 mw mass flow in ltr (cold s.) 2299.89 kg/s mass flow in com. no.2 430.1 kg/s mass flow ratio 0.842 net power 201.69 mw mass flow 2730.0 kg/s pressure ratio 2.8 pressure no.2 34.0 mpa table 3. result of gfr re-compression cycle. the calculation of s-co2 cycle for demo2 was performed for two cycle. the first cycle is the recompression cycle and the second cycle is preheating cycle. the calculations were performed with the parameters from the table 2. the cycles were optimized for the best results. the result of the demo2 recompression cycle is shown in the table 4. the layout of heat sources corresponds to the figure 2. the heat source h is split into two heat sources arranged serially. according to the table 4, the total net power of the re-compression cycle is 1251 mw. the cycle efficiency is 30.45 %. however, the mass flow for this case is very high and achieves of 20 250 kg/s. cycle efficiency 30.45 % turbine power output 2013.02 mw compressor no.1 in. power 719.62 mw compressor no.2 in. power 42.41 mw compressor input power 762.03 mw added heat 4108.83 mw removed heat 2857.84 mw regenerative heat 8340.95 mw mass flow in ltr (cold s.) 19 589.33 kg/s mass flow in com. no.2 660.66 kg/s mass flow ratio 0.967 net power 1250.98 mw mass flow 20 250.0 kg/s pressure ratio 2.6 pressure no.2 34.0 mpa table 4. result of demo2 re-compression cycle. the figure 8 shows the t-s diagram of the demo2 re-compression cycle. the result of the demo2 pre110 vol. 4/2016 comparison of s-co2 power cycles for nuclear energy figure 8. t-s diagram of the demo2 re-compression cycle. figure 9. t-s diagram of the demo2 preheating cycle. heating cycle is shown in the table 5 and the t-s diagram of the cycle is shown in the figure 9. the layout of heat sources corresponds to the figure 6. the heater h1 was used for high-grade primary heat, the heater h2 was used for low-grade secondary heat. according to the table 5, the total net power of preheating cycle cycle is 1145 mw. the cycle efficiency is 27.87 %, and the mass flow for the preheating cycle is 14 040 kg/s. 111 l. vesely, v. dostal, s. entler acta polytechnica ctu proceedings cycle efficiency 27.87 % turbine power output 1727.14 mw compressor input power 582.005 mw added heat 4108.54 mw removed heat 2963.41 mw regenerative heat 4125.79 mw mass flow in hex 13 338.0 kg/s mass flow in h2 702.0 kg/s mass flow ratio 0.95 net power 1145.13 mw mass flow 14 040 kg/s pressure ratio 3.4 pressure no.2 28 mpa table 5. result of the demo2 preheating cycle. 7. conclusion the results from the table 3 represent the suitable results for the gfr designing with re-compression cycle. improvement of the parameters is possible by future research. the results of the s-co2 cycle for the fusion reactor show the effect of the use of multiple heat sources. the re-compression cycle is better than the preheating cycle according to the table 4 and table 5. however, it is obvious that this advantage is valid only for the net power. the critical disadvantage of the recompression cycle is the mass flow of about 20 000 kg/s. the mass flow of the preheating cycle is lower in comparison with the re-compression cycle. designing and optimization of the s-co2 cycles are of great importance for fusion energy. the preheating cycle is a convenient cycle for utilization of low potential heat, and such cycles are suitable for exploitation of the multiple heat sources of the fusion power reactor like analyzed demo2. further research of the s-co2 power cycles for the fusion power reactors will be focused on the detailed comparison of the s-co2 power cycle and on the development of the new modifications of the s-co2 power cycle, which will take into account the fusion reactor multiple heat source design. references [1] g. angelino. carbon dioxide condensation cycles for power production. j eng power 90(3):287–295, 1968. doi:10.1115/1.3609190. [2] v. dostal, m. driscoll, p. hejzlar. a super critical carbon dioxide cycle for next generation nuclear reactors, vol. mit-anp-tr-100 of advanced nuclear power program. mit center for advanced nuclear energy systems, 2004. [3] v. ladislav, d. vaclav, b. ondrej, n. vaclav. pinch point analysis of heat exchangers for supercritical carbon dioxide with gaseous admixtures in ccs systems. energy procedia 86:489–499, 2016. doi:http://dx.doi.org/10.1016/j.egypro.2016.01.050. [4] e. j. parma, s. a. wright, m. e. vernon, et al. supercritical co2 direct cycle gas fast reactor (sc-gfr) concept. tech. rep. sand2011-2525, sandia national laboratories, 2011. [5] e. a. harvego, m. g. mckellar. optimization and comparison of direct and indirect supercritical carbon dioxide power plant cycles for nuclear applications. in asme 2011 international mechanical engineering congress and exposition, pp. 75–81. 2011. doi:10.1115/imece2011-63073. [6] b. halimi, k. y. suh. computational analysis of supercritical {co2} brayton cycle power conversion system for fusion reactor. energy conversion and management 63:38–43, 2012. doi:http://dx.doi.org/10.1016/j.enconman.2012.01.028. [7] j. i. linares, l. e. herranz, i. fernández, et al. supercritical co2 brayton power cycles for demo fusion reactor based on helium cooled lithium lead blanket. applied thermal engineering 76:123–133, 2015. [8] g. federici, r. kemp, d. ward, et al. overview of eu demo design and r&d activities. fusion engineering and design 89(7–8):882–889, 2014. doi:http://dx.doi.org/10.1016/j.fusengdes.2014.01.070. [9] r. kemp. demo2 reference design, 2015. eurofusion idm eu_d_2lcbvu. [10] e. w. lemmon, m. l. huber, m. o. mclinden. nist reference fluid thermodynamic and transport properties – refprop. national institute of standards and technology, 2013. version 9.1. 112 http://dx.doi.org/10.1115/1.3609190 http://dx.doi.org/http://dx.doi.org/10.1016/j.egypro.2016.01.050 http://dx.doi.org/10.1115/imece2011-63073 http://dx.doi.org/http://dx.doi.org/10.1016/j.enconman.2012.01.028 http://dx.doi.org/http://dx.doi.org/10.1016/j.fusengdes.2014.01.070 acta polytechnica ctu proceedings 4:107–112, 2016 1 introduction 2 advantages and disadvantages of s-co2 cycles 3 description of gas cycles 4 the gfr re-compressing cycle 5 the s-co2 cycle for demo2 fusion reactor 6 result of gfr reactor and demo2 fusion reactor 7 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0004 acta polytechnica ctu proceedings 5:4–11, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app on efficient operational concept of future high-speed railway in the czech republic michal drábek department of logistics and management of transport, faculty of transportation sciences, ctu in prague, czech republic correspondence: xdrabek@fd.cvut.cz abstract. the aim of this paper is to elaborate a layout of the first operational concept of rapid services with 1 hour system travel time between praha and brno. two basic methods are used – integrated periodic timetable (periodic rendezvous of all services in ipt-nodes) and operational concept economy approach, as defined below by the author. in this paper, three recent high-speed railway concepts for the future so-called rapid services network of the czech republic are followed-up. the first one is an operational traffic planning study by kalčík, janoš et al. on behalf of czech ministry of transport from 2010. the second one is the high-speed railway promoting book high speed rail even in the czech republic by šlegr et al. from 2012, with likely the most detailed concept of rapid services network. the third one is a paper on progress of the official spatial-technical studies for some future czech high-speed lines by šulc from 2014. the importance of achievement of 1 hour travel time between the largest agglomerations is briefly presented. the presented methodological approach, although soft and manager-oriented, comprises some firm principles: segmentation of high-speed train offer, so that more expensive rolling stock is not wasted by operation on long conventional line sections, consideration of system travel times for efficient rolling stock circuit, restriction of need for links from high-speed to conventional lines, and utilization of high-speed lines as a "rail highway". this approach is intended to be particularized iteratively, with every application. so, in this paper, first version of operational concept economy approach is introduced. the key idea is that passengers should be offered such travel times and service intervals (headways) and such number of direct services, which are adequate to their potential demand, but as much synergistic effect as possible should be strived to be achieved for every proposed construction (new or modernized one). such approach goes towards economic efficiency, which is crucial indicator for political decision necessary for building, let alone eu co-funding of the construction. experience shows that in many czech feasibility studies, achievement of sufficient economic efficiency was the most complicated part of the study. results show that an efficient operational concept can be designed not at the expense of runtimes between the largest cities. keywords: high-speed railway, high-speed line, rapid services, integrated periodic timetable . 1. introduction the aim of this paper is to elaborate a layout of the first operational concept of rapid services with 1 hour system travel time between praha and brno. two basic principles are used – integrated periodic timetable, and operational concept economy approach, as defined below by the author. directive 2008/57/ec of the european parliament and of the council of 17 june 2008 on the interoperability of the rail system within the community [1] defines in annex ii infrastructure and rolling stock as structural subsystems of the rail system, and traffic operation and management as a functional subsystem of the rail system. in traffic operation and management subsystem, traffic planning is also included. this paper is organized as follows. after list of abbreviations and list of largest czech cities, recent concepts of high-speed lines network for rapid services, in the context of recent ten-t regulation [2] are introduced. then, significance of 1 hour travel time is summarized, based mostly on revealedpreference data. introduction of used methods follows: integrated periodic timetable and newly developed soft operational concept economy approach. further, scheme of the resulting operational concept is presented, commented and discussed. the paper is closed by conclusion, including recommendation for further studies in this field. for the sake of brevity and clarity, following abbreviations will be used in this paper: • emu – electric multiple unit • hs – high-speed (railway) • hsl – high-speed line • ipt – integrated periodic timetable • put line – public transport line (not to be mistaken with railway line as a built structure) – a group of services which operate all day in regular period, and serve particular sequence of stations and stops • rs – rapid services (czech system of hsls and follow-up uprgraded or new lines, in czech: rychlá spojení) 4 http://dx.doi.org/10.14311/app.2016.5.0004 http://ojs.cvut.cz/ojs/index.php/app vol. 5/2016 on efficient operational concept • sždc – railway infrastructure administration, state organization – czech state railway infrastructure manager ( in czech: správa železniční dopravní cesty ) • ten-t – trans-european transport networks • tsi – technical specification of interoperability • uic – international union of railways 2. list of largest czech cities table 1 displays list of all czech cities over 40,000 inhabitants. these cities can be divided into four subgroups based on their population, as of 1st january 2015. regional capitals are emphasized. note that population of ostrava agglomeration is comparable with praha. these cities build up five largest agglomerations: praha, ostrava region, brno, northwestern bohemia (from děčín to chomutov) and central moravia. 3. ten-t context and recent concepts of hsls for rapid services most of the intended czech hsls are part of ten-t network, but only three of them are involved into the core ten-t network: hsls praha – ústí nad labem and brno – břeclav, and modernization of the line brno – přerov. for the hsl brno – břeclav, there are several alternative directions southward of vranovice. for the hsl praha – wroclaw (in comprehensive network), there are two alternative directions: via liberec or via hradec králové. the author has followed up from three main sources. each one has been carried out for different purpose. the study carried out by kalčík, janoš et al. (including the author) [4] contains, to the author’s knowledge, the only model timetable for the whole rs network up to now. approximate routes for hsls were defined, and, based on their parameters and on available hs rolling stock parameters, model timetables were constructed using timetabling software fbs. one of the results is a netgraph (scheme of put lines with arrival and departure minutes in node stations). šlegr et al. [5] have written and published the book high speed rail even in the czech republic with many hsl drawings in the map. the book introduces european and world best practice in hs railway and, in a detailed manner, presents potential benefits of the rs system for every czech region. as a result, the most detailed concept of rs network (although not officially approved), with qualified estimation of section runtimes, is introduced there. šulc [6] has presented recent progress in five spatial-technical studies of hsls, carried out on behalf on sždc. the aim of these studies is to find suitable hsl routes, so that necessary space for their future construction can be legally protected. for this purpose, simplified timetable layout and cost estimation were carried out. other hsls were already subject of more detailed feasibility studies: praha – liberec and brno – přerov. the latter study recommended the doubling and modernization of present single-track line up to 200 km/h as the best one, although the assessed line is located directly in the corridor brno – ostrava. uic [7] records that hsls for the maximum speed of 350 km/h are in operation in china (1,028 km), and they are planned in portugal (206 km). in the united kingdom, hsls for the maximum speed of 360 km/h (543 km) are planned. thus, such maximum speed is nothing uncommon for the recently built and planned hsls. figure 1 shows the maximalist future czech hsl network as a backbone of rapid services. it does not include possible local new lines or line modernizations. the scheme was created as follows. for each hsl, always the higher maximum speed from šlegr et al. [5] and šulc [6] was chosen. inclusion of each hsl (or its section) into ten-t core, or comprehensive network based on [2] was marked. the actual phase of project preparation (based on šulc [6] and to the author’s knowledge) was also marked. from the possible links to conventional railway network, only the most significant ones were displayed. the stations or stops, which are intended to be built directly on hsls were adopted from šulc [6]. international hsls are strongly dependent on the will of both neighbouring states: the hsl from ústí nad labem to dresden is supported by czech ministry of transport, as well as the government of saxony (although it requires a long tunnel under the mountains), but the hsl from plzeň to bavaria is not quite a priority for the government of bavaria, despite its indisputable european significance. 4. one hour as a breakthrough in travel time the practice has proven that, when the public transport travel time changes to 1 hour or less, and if this travel time is competitive to car, public transport ridership considerably increases, and some people can even commute daily. if the present travel time by cars between two largest czech cities is approximately 2 hours, only 1 hour is fully competitive, because the travel time to and from the station must be added. in switzerland, new line for 200 km/h was built to achieve travel time of 56 min between zurich and berne (swiss largest city and capital). this line was not an isolated project, but a part of targeted investments for implementing bahn 2000 concept. its main aim was to achieve necessary system travel times for the ipt. this new line, as a kind of swiss ”rail highway”, serves also for many other long-distance put lines. the modal-split of railway for the journeys zurich – berne has achieved 88%, for the journeys zurich – st. gallen (1:02 h travel time) it was 73% (voev utp, 2010 [8]). janoš and baudyš [9] state that improvements in 5 michal drábek acta polytechnica ctu proceedings city population city population praha 1,259,079 zlín 75,112 brno 377,440 havířov 75,049 ostrava 294,200 kladno 68,552 most 67,089 opava 57,772 frýdek místek 56,945 karviná 55,985 city population city population plzeň 169,033 jihlava 50,521 liberec 102,562 teplice 50,079 olomouc 99,809 děčín 49,833 ústí nad labem 93,409 karlovy vary 49,781 české budějovice 93,285 chomutov 48,913 hradec králové 92,808 jablonec nad nisou 45,594 pardubice 89,693 mladá boleslav 44,318 přerov 44,278 prostějov 44,094 table 1. list of largest czech cities. data from czech statistical office [3]. figure 1. maximalist czech hsl network 6 vol. 5/2016 on efficient operational concept timetable offer (more frequent and more periodic service) in long-distance railway in the czech republic have led to increase of number of passengers. ”between 2003 and 2005, incomes increased by 7.86%, and number of passengers by 3.63%. however, in some relations, thanks to new supply, the increase of passengers was 40 to 50%. ipt is being further developed. thanks optimal train connections in junctions and synergic effect of linked supply number of passengers raises by 2-10% per year even on lines with no supply change.” the raise of service for timetable 2004/05 was 7%, compared to timetable 2003/04, when new offer (supply) concept was introduced. vichta [10] has also shown figures that justified timetable offer improvement for large enough cities (ústí nad labem and teplice). since 2003/04 timetable, two segments of periodic service between ústí and praha were introduced: a non-stop fast train with travel time slightly above 1 hour and a 120 min interval, and fast train with several stops hourly. since 2008, a non-stop segment has been operated hourly, and every second service connected teplice directly with praha. compared with train ridership from 2002 (100%), the train ridership between ústí and praha grew to 152% in 2006, and to 325% in 2008. compared to the same year, the train ridership between teplice and praha grew to 228% in 2006 and to 369% in 2008. in both cases, only relative changes can be published, to avoid publishing of czech railways’ business sensitive data. kordis jmk, an integrated transport system coordinator of south moravia region, has carried out a stated-preference study of potential ridership of hsl between praha and brno [11]. the study has proved that about 30,000 passengers daily would use the new hsl, majority of them instead of cars. the potential ridership varied depending on train travel time and ticket price. the results for 45 min travel time were only slightly better than the ones for 60 min. the acceptable ticket price was up to 299 czk. for 300 czk and more, the potential ridership fell to approximately 18,000 passengers. a travel time between praha and brno up to 60 min is also supported officially, at least by some departments of czech ministry of transport and sždc – see šulc, page 14 [6] and kušnír and ilík, page 13 [12]. the graph in the latter presentation shows clear intention to ensure better connection not only for the largest agglomerations, but also for the regions, so that their development can be stimulated. however, a more detailed study on potential rs ridership, which should include a transport demand forecasting model and more detailed economic calculations, has to be carried out. sždc has called a bid for so called ”rapid services – study of opportunities”, that had to be cancelled, because only one eligible applicant remained. recently, a new bid for the study is being in progress. 5. methods 5.1. principle of ipt ipt is a special case of periodic timetable, which ensures connections between various put lines in the whole network. for this effect, the following requirements must be fulfilled. all services operate in put lines, in unified interval which is equal to 2k-multiple of basic period (60 min as a rule), where k is an integer. in every put line, services from opposite directions meet each other at the same time (symmetry time). this time repeats after half an interval. should services of two put lines enable mutual connections with an equal changing time for each direction, these two put lines must have equal symmetry time. in european long-distance railways, symmetry time slightly before the minute 00 is common. this is called zero symmetry axis. in practice, symmetry times in the minute 57 to 01 are used, which, in the case of hourly interval, implies another symmetry time in the minute 30. for the 30 min interval, additional symmetry times in the minutes 15 and 45 occur. an ipt-node is a railway station where connections with services from other put lines can be ensured, because the symmetry time is reached there. to achieve connections in every ipt-node, system travel time between them must be ensured. it is an integer multiple of half period. it consists of sum of regular runtimes between two iptnodes, sum of dwell times and waiting times between them, and of proportional part of dwell, changing or waiting times in mentioned ipt-nodes (it depends which time is limiting in each particular case) [13]. it is obvious that two put lines with equal runtimes, which operate on long enough common line section and do not need to be connected for the passenger transfer, can be interposed into a half interval. this interposition is used for more attractive timetable offer. on the other hand, if the transfer is desirable, the trains of two put lines run as close as possible after each other (so that they reach the same ipt-node, e.g., in berne). 5.2. operational concept economy approach the presented methodological approach, although soft and manager-oriented, comprises some firm principles, which will be explained as follows. this approach is intended to be particularized iteratively, with every application. so, in this paper, first version of operational concept economy approach is introduced. the key idea is that passengers should be offered such travel times and service intervals (headways) and such number of direct services, which are adequate to their potential demand, but as much synergistic effect as possible should be strived to be achieved for every proposed construction (new or modernized one). such approach goes towards economic efficiency, which is a crucial indicator for political decision necessary for building, let alone eu co-funding of the construction. 7 michal drábek acta polytechnica ctu proceedings experience shows that in many czech feasibility studies, achievement of sufficient economic efficiency was the most complicated part of the study. travel times should correspond to system travel times according to ipt rules. of course, system travel time is always considerably larger than a corresponding runtime, because of times for changing between services in the ipt nodes, and necessary time supplement for the sake of timetable stability. because of high-speed railway dynamics, such supplement should be larger here than by the conventional railway. runtime calculations are not subject of this paper, but it is obvious, that for approx. 200 km distant cities praha and brno and for system travel time 60 min between them, maximum speed of 200 km/h is totally insufficient. so, higher speed, e.g. 350 km/h, is considered. moreover, in city centres, only lower speed is possible due to densely built-up areas and probably one more stop in praha before the centre (zahradní město) [6]. if a new hsl is justified by the potential demand, it should be at the same time used by as many trains as possible, to make faster as many connections as possible. thus, at least central parts of the hs network should serve as ”rail highways” for fast enough long-distance, or even fast regional, put lines. this idea leads to requirements for junctions to connect a hsl with conventional rail network. a recent concept of rs of czech ministry of transport [12] corresponds to such idea, because of ensuring fast railway service also for czech regions. but, because of economy, there should be less junctions with as much put lines using one junction as reasonable. as a rule of thumb, every junction from hs line should be used by minimum 1 pair of trains hourly. the design of put lines themselves should also be in compliance with economy. for high-speed railway, a hierarchy in the form of following segments of offer can be defined: • fastest put lines (hs-a) for the speed up to 350 km/h, with dedicated emus. in figure 2, such put lines are marked in red colour. • long-distance put lines (hs-b1) which operate on long sections of fast high-speed lines, so the rolling stock should be able to run 230 to 250 km/h to avoid frequent overtaking by trains of higher segment. rolling stock can vary from emu to classical trainset, or push-pull units hauled by a locomotive (e.g. railjet). in figure 2, such put lines are marked in violet colour, and in pink colour, in the case of tilting trains. • long-distance put lines (hs-b2) which operate on short sections of fast high-speed lines, or they are overtaken by hs-a trains. thus, for the rolling stock the maximum speed of 200 km/h is sufficient. rolling stock can vary from emu to classical trainset. in figure 2, such put lines are marked in blue colour. • fast regional put lines (hs-c), which can achieve significantly lower travel time to the agglomeration centre by using of hsl section. the rolling stock should run at least 160 km/h and be interoperable with tsis required for high-speed railway (for the other segments, this is evident). for maintaining reasonable complexity, this segment is neglected in this paper. for the sake of economy, the higher segment, the less trainsets should be required. this aim can be achieved by two means: as little number of put lines as reasonable, and as little number of trainsets per put line as reasonable. but, on the other hand, appropriate service interval, and some reserve trainsets should be ensured. the basic service interval for each put line is considered 60 min. the only exception is international put line warsaw – vienna, which is of international more than national significance. the basic service interval of 30 min (of a single put line or an interposition of them) is considered for connection of all regional centres above 90 000 inhabitants with praha, brno or ostrava, if the system travel time to the particular agglomeration is 1 hour or lower. the hs-a segment should end in cities, which are large enough to justify the most expensive and maintenance sensitive rolling stock. for the hs-a segment between praha and brno, a basic interval 15 min is considered. in some studies, even an interval 10 min is discussed, but it would complicate operation of hs-b trains on the same hsl. so, the longest possible trains (400 m), and double-decker trains (such as tgv duplex) are considered as mostly desirable. the number of four basic hs-a put lines (each one with 60 min interval) should not be exceeded. thus, every hs-a train is economically justified because of service between praha and brno. because of ”biggest possible” trains, a 30 min period of hs-a segment between praha and ostrava is likely to be sufficient. it is obvious that vast majority of the demand would be national, mostly between large cities. so, for the sake of economy, a standard hs-a trainset consisting of two half-trains (such as ice 3, but double-decker) is further proposed. this measure would enable eventual disconnection of one trainset in the station, where many passengers leave and only little number of passengers board. such disconnection enables also direct service to two various destinations. as a rule of thumb for the czech republic, middle-sized cities with 30,000 or more inhabitants can be considered as appropriate ends of hs-a put lines. it is, however, important that a runtime from regional centre to such city must not exceed approximately 25 minutes. thus, service for such middle-sized city in 60 min interval would cost only 1 trainset extra (a half-train only). for the hs-b segment, the choice of rolling stock depends on timetable of hsl section used (required maximum speed), and on justification of tilting technology (a long enough conventional line with many curves, and potential time savings related to the population served). the put lines should be composed with respect to equal rolling stock requirements on both sides of a metropolis (for 8 vol. 5/2016 on efficient operational concept the czech republic, this means in practice praha and brno). for the hs-c segment, runtime savings should justify extra costs for ”high-speed-compatible” technical equipment of the rolling stock. thus, such trains should use high-speed line sections both before and after stop in the agglomeration centre. for the feasibility of the proposed operational concept, electrification and upgrading of several mainlines is supposed. further, a new line from třebíč to moravské budějovice is required, to enable direct connection of jihlava, třebíč and znojmo. for the sake of reasonable scope, timetables are not constructed. it is supposed, that in every ipt-node, slower trains leave after faster trains, if they run on the same line. thanks to the symmetry of ipt, arrival sequence is symmetric (slower trains arrive before faster ones). the system travel times are estimated on the basis of kalčík, janoš et al. [4], and lengthened due to additional stops if needed (by qualified estimation). 6. results and discussion the scheme of put lines, ipt-nodes and system travel times between large cities is displayed in figure 2. an interposition into half interval (30 min) is marked by two parallel lines close together. for the hs-a segment, the interpositions into 30 min and 15 min intervals have resulted in the fact that either plzeň or ústí nad labem, but not both cities, can be connected directly with ostrava. considering importance of direct connection of berlin and vienna via praha for the czech republic, ostrava should be connected directly with plzeň rather than ústí. thus, two bundles of hs-a put lines emerged: • berlin – dresden – praha – brno – břeclav – vienna/bratislava – budapest and litvínov/děčín – ústí nad labem – praha – brno – zlín • munich (or nuremberg) – plzeň/plzeň – praha – brno – ostrava – petrovice u karviné – warsaw/opava and plzeň – praha – brno – ostrava – havířov – žilina the sign ”/” stands for coupling and decoupling of half-trains. for international directions, full trains were only considered for berlin (because of national demand between dresden and berlin) and žilina, because of traditional ”czechoslovak” passenger demand (e.g. students, workers or tourists). from plzeň, only a half-train proceeds to bavaria. in břeclav, decoupling enables direct service from praha both to vienna, and bratislava (and budapest). the put line, which ends in ústí nad labem, can proceed both to děčín and teplice, thanks to decoupling. if the line to litvínov would be upgraded, a half-train can end there. so, service of both branches would cost only 1 half-train extra per branch. from ostrava, only a half-train would proceed to warsaw. thus, the second half-train can proceed to opava, and even with present runtime it would cost only 1 half-train extra. hs-b1 segment is proposed mostly in moravia for the connection of central moravia between brno and ostrava. the reason for faster trains is the absence of stops directly on the hsl brno – ostrava. the absence of clear centre resulted in necessity of more put lines. olomouc is connected with brno by a non-stop service every 30 min. every second train proceeds north-westward to connect mohelnice, zábřeh and šumperk directly with brno. due to longer runtime, two extra trains are required. olomouc is also connected with ostrava every 30 min. one put line proceeds as an express segment to pardubice and hradec králové, the second one as a slower inter-regional train to pardubice. from pardubice to praha, the same put line changes to a non-stop segment. between praha and olomouc, an interposition into 30 min interval is proposed. the interposing put line proceeds to vsetín. because the last put line uses hsl only between praha and pardubice, a trainset for 200 km/h could be appropriate. přerov is connected hourly both with brno and ostrava. behind přerov, both put lines proceed as a conventional fast train. hradec králové is connected with praha only hourly. the supplementary connection can be ensured by transfer in pardubice. the put line from hradec králové is proposed to be composed of two emus, because of further decoupling, so that both trutnov and náchod can be connected directly with praha. liberec is connected with praha every 30 min, with only intermediate stop in mladá boleslav. jablonec nad nisou is connected with liberec (and hence with praha) by suburban trains. the same case is kladno with praha (but there are supposed fast regional trains, running directly to praha main station). north-western bohemia is connected with praha every 30 min, with likely stop in louny (for the connection of this peripheral region) between most and praha. one of two put lines proceeds to karlovy vary and cheb, and, because of the terrain, tilting train operation can save considerable amount of runtime. both mentioned put lines proceed from praha to jihlava, but with various stopping. the tilting train as hs-b1 segment runs nonstop, and proceeds to třebíč and znojmo (possible proceeding to vienna is worth consideration), where tilting is supposed to save the runtime. another put line stops in benešov, and on the hsl in vlašim and vysočina (a stop between pelhřimov and humpolec, linked to bus and park+ride). overtaking by hs-a trains in some of these stations is very likely – that is why hs-b2 segment is sufficient. from jihlava, this put line proceeds to havlíčkův brod and žďár nad sázavou. instead of separate link from hsl to havlíčkův brod, a thorough modernization and doubling of the line jihlava – havlíčkův brod is proposed here. jihlava is connected with brno every 30 min, with intermediate stop in velké meziříčí rs, most likely with overtaking by hs-a segment at the same time. thus, hs-b2 segment seems to be sufficient. from jihlava, one put line would proceed to havlíčkův brod, and the second one to české budějovice. modernization of 9 michal drábek acta polytechnica ctu proceedings figure 2. proposed operational concept for hsl network of rs. the line jihlava – veselí nad lužnicí and a new tunnel ševětín – nemanice are supposed to be comlpeted. the last put line of hs-b2 segment is děčín – ústí nad labem – litoměřice – roudnice nad labem rs – praha – tábor – české budějovice. the corridor to české budějovice is supposed to be finished. one of the most important results is a reduction of links from hsls to conventional lines outside large nodes. for long-distance put lines (hs-b1 and hs-b2) there have remained: litoměřice, benešov u prahy (twice), jihlava (twice), olomouc (twice) and přerov (twice) – in total 9 links. it is desirable to try to design links near olomouc and přerov together as well as to design single-track links, if they would be used by only one pair of trains hourly. of course, the proposed operational concept implies plenty of alternatives. for instance, a 30 min interval praha – hradec králové, or another put line layout for the connection of jihlava. if the put line interval is 60 min, the timetable should be constructed so that additional peak trains (which leads to interposition into 30 min interval) can be added without any conflict. the probably biggest disadvantage of the rs concept used by the author is that pardubice cannot be connected with brno within 1 hour system travel time. the author is aware of alternative concepts of hsl praha – brno via pardubice. but, given that system travel time 1 hour between praha and brno is not quite easy to achieve even via jihlava, the longer hsl via pardubice would hardly enable such time, even with a non-stop train running up to 350 km/h. 7. conclusion the presented operational concept is only a brief layout, which served as a test of operational concept economy approach, and a feedback for future making this approach more accurate. the author believes than his layout would also contribute to constructive discussion on future czech rs system, and on the role of hsls in it. the author considers as an important contribution in this paper, that there were shown close economic interdependencies between operational concept, rolling stock and infrastructure, and the operational concept was designed on the basis of such interdependencies, towards economic efficiency, but not at the expense of runtimes between the largest cities. the czech rs network always will be an integral part of both european and national transport system. thus, its layout and construction will be always subject of political decisions. the presented operational concept has clearly demonstrated many possible alternative solutions. although detailed studies will calculate which alternative brings higher benefits for lower costs, the most likely alternative to be constructed will be, in the 10 vol. 5/2016 on efficient operational concept author’s opinion, a result of political compromise or trade-off. but, even by choosing any such sub-optimal alternative the decision makers should keep in mind than it should enable efficient allocation of resources. there is always a high risk not only of construction of excessive infrastructure, but also of inefficient operational concept (too many put lines for too low demand). thus, every alternative, before its detailed economic assessment, has to be designed cautiously to minimize such risk. the compromises are inevitable, but they should not generate disproportionate extra costs. acknowledgements the author would hereby like to propose a vote of thanks to mr. petr šlegr for his long-standing promotion of highspeed railway in the czech republic, which he has crowned by editing and publishing book high speed rail even in the czech republic [5]. the author would hereby like to propose a vote of thanks to dr. vít janoš, a supervisor of his doctoral thesis, for initial layout of an operational concept for rs [4], that the author was glad to follow up. the author would hereby like to propose a vote of thanks to mr. jiří kalčík, a freelance road and railway designer, for searching for efficient high-speed lines, both in official studies (e.g. [4]), and in his spare time [5]. so, he has contributed significantly to the concept mentioned above. all engineers enabled the author to follow their ideas from timetabling viewpoint – for instance, in this paper. the author would hereby also like to propose a vote of thanks to sždc as a former author’s employer, for the occasion to assess five spatial-technical studies of hsls simultaneously from timetabling viewpoint (in summer 2013). this task has made the author to create his own view on efficient operational concept for the future czech rs network. his interest in this field has endured up to now, and has resulted in writing this paper. the author hopes it will not remain his last and the most detailed work in this field. references [1] european commission. directive 2008/57/ec of the european parliament and of the council of 17 june 2008 on the interoperability of the rail system within the community [online], 2011. european union ,brussels, http://eur-lex.europa.eu/legal-content/en/txt/ pdf/?uri=celex:02008l0057-20110322&from=en. [2] european commission. regulation (eu) no 1315/2013 of the european parliament and of the council of 11 december 2013 on union guidelines for the development of the trans-european transport network and repealing decision no 661/2010/eu [online]. 2013. european union ,brussels, http://eur-lex.europa.eu/legal-content/en/all/ ?uri=uriserv:oj.l_.2013.348.01.0001.01.eng. [3] czech statistical office. populations of municipalities [online], 2015. [2016-04-30], https://www.czso.cz/ csu/czso/pocet-obyvatel-v-obcich-k-112015. [4] j. kalčík, v. janoš. operational traffic planning solution for backbone network of high-speed lines, study on behalf of czech ministry of transport. chrást u plzně: projektové středisko kalčík, czech republic, 2010. unpublished. [5] p. šlegr et al. high speed rail even in the czech republic, 2012. editors: p. šlegr, m. robeš, m. drábek, m. stach; isbn 978-80-905005-0-1. [6] j. šulc. actual progress of preparation of rapid services in sždc. in: czech raildays: building of high-speed rail system in the czech republic. separate article [online]. praha: m-presse plus s.r.o., 2014. in czech, http://www.railvolution.net/ czechraildays/2014/buletin_szdc.pdf. [7] uic. high speed lines in the world [online], 2014. paris: uic, http://old.uic.org/img/pdf/20140901_ high_speed_lines_in_the_world.pdf. [8] voev utp: fact and arguments in favour of swiss public transport. p. 46. bern: voev utp, 2010. in german. [9] v. janoš, k. baudyš. railway timetable in czech republic, in systemy transportowe teoria i praktyka, 2009. isbn 978-83-926923-1-7. [10] f. vichta. requirements for public transport system in the czech republic, keynote speech, in future of railway passenger transport in the czech republic, czech raildays [online], 2011. ostrava: railvolution, [2011-06-14], in czech, http://www.railvolution.net/ czechraildays/2011/seminare/1.3_md_vichta.pdf. [11] railhuc project. final results [online], 2014. brno: kordis jmk, isbn 978-80-260-7787-9, http: //www.railhuc.eu/download/category/126-book2. [12] j. kušnír, j. ilík. railway in 2030, keynote speech, in future of railway passenger transport in the czech republic, czech raildays [online], 2013. ostrava: czech ministry of transport, in czech., http://www.railvolution.net/czechraildays/2013/ seminare/konference-kusnir.pdf. [13] m. drábek. periodic freight train paths in network, doctoral thesis, 2014. ctu in prague, http://takt.fd.cvut.cz/cargo/drabek_thesis.pdf. 11 http://eur-lex.europa.eu/legal-content/en/txt/pdf/?uri=celex:02008l0057-20110322&from=en http://eur-lex.europa.eu/legal-content/en/txt/pdf/?uri=celex:02008l0057-20110322&from=en http://eur-lex.europa.eu/legal-content/en/all/?uri=uriserv:oj.l_.2013.348.01.0001.01.eng http://eur-lex.europa.eu/legal-content/en/all/?uri=uriserv:oj.l_.2013.348.01.0001.01.eng https://www.czso.cz/csu/czso/pocet-obyvatel-v-obcich-k-112015 https://www.czso.cz/csu/czso/pocet-obyvatel-v-obcich-k-112015 http://www.railvolution.net/czechraildays/2014/buletin_szdc.pdf http://www.railvolution.net/czechraildays/2014/buletin_szdc.pdf http://old.uic.org/img/pdf/20140901_high_speed_lines_in_the_world.pdf http://old.uic.org/img/pdf/20140901_high_speed_lines_in_the_world.pdf http://www.railvolution.net/czechraildays/2011/seminare/1.3_md_vichta.pdf http://www.railvolution.net/czechraildays/2011/seminare/1.3_md_vichta.pdf http://www.railhuc.eu/download/category/126-book2 http://www.railhuc.eu/download/category/126-book2 http://www.railvolution.net/czechraildays/2013/seminare/konference-kusnir.pdf http://www.railvolution.net/czechraildays/2013/seminare/konference-kusnir.pdf http://takt.fd.cvut.cz/cargo/drabek_thesis.pdf acta polytechnica ctu proceedings 5:4–11, 2016 1 introduction 2 list of largest czech cities 3 ten-t context and recent concepts of hsls for rapid services 4 one hour as a breakthrough in travel time 5 methods 5.1 principle of ipt 5.2 operational concept economy approach 6 results and discussion 7 conclusion acknowledgements references 277 acta polytechnica ctu proceedings 2(1): 277–281, 2015 277 doi: 10.14311/app.2015.02.0277 what powers the 2006 outburst of the symbiotic star bf cygni? a. skopal1, m. sekeráš1, n. a. tomov2, m. t. tomova2, t. n. tarasova3, m. wolf4 1astronomical institute of the slovak academy of sciences, 059 60 tatranská lomnica, slovakia 2institute of astronomy and nao, bulgarian academy of sciences, 4700 smolyan, bulgaria 3crimean astrophysical observatory, 298409 nauchny, crimea, russia 4astronomical institute, charles university prague, 180 00 praha 8, v holešovičkách 2, czech republic corresponding author: skopal@ta3.sk abstract bf cygni is a classical symbiotic binary. its optical light curve occasionally shows outbursts of the z and-type, whose nature is not well understood. during the 2006 august, bf cyg underwent the recent outburst, and continues its active phase to the present. the aim of this contribution is to determine the fundamental parameters of the hot component in the binary during the active phase. for this purpose we used a highand low-resolution optical spectroscopy and the multicolour ubv rcic photometry. our photometric monitoring revealed that a high level of the star’s brightness lasts for unusually long time of > 7 years. a sharp violet-shifted absorption component and broad emission wings in the hα profile developed during the whole active phase. from 2009, our spectra revealed a bipolar ejection from the white dwarf (wd). modelling the spectral energy distribution (sed) of the low-resolution spectra showed simultaneous presence of a warm (< 10 000 k) disk-like pseudophotosphere and a strong nebular component of radiation (emission measure of ∼ 1061 cm−3). the luminosity of the hot active object was estimated to > 5 − 8 × 103 l�. such high luminosity, sustained for the time of years, can be understood as a result of an enhanced transient accretion rate throughout a large disk, leading also to formation of collimated ejection from the wd. keywords: binaries: symbiotic optical spectroscopy photometry individual: bf cyg. 1 introduction symbiotic stars (sss) are the largest interacting binary systems with known orbital periods in order of years. they consist of a cool giant and a wd accreting from the giant’s wind. accretion process heats up the wd to 1 − 2 × 105 k and makes it as luminous as a few times 103 l�, whose photons ionize a large fraction of the neutral giant’s wind, giving rise to nebular emission. as a result the spectrum of sss consists of three basic components of radiation – two stellar and one nebular. if a symbiotic system releases its energy approximately at a constant rate and the temperature, it conforms the so-called quiescent phase. the stage, when the system brightens up in the optical by a few magnitudes and/or shows signatures of a mass-outflow, is named an active phase. bf cyg is an eclipsing symbiotic binary with an orbital period of 757.2 d (e.g. pucinskas 1970; fekel et al. 2001). the binary consists of a late-type m5 iii giant (mürset & schmid, 1999) and a hot luminous compact object (mikolajewska et al., 1989). its eclipsing nature was revealed by optical photometry of skopal (1992). historical light curve of bf cyg is characterized by a aaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaa slow, symbiotic-nova-type outburst (1895–1960), with superposed eruptions of the z and-type (1920, 1989, 2006) and short-term flares (e.g. skopal et al., 1997; leibowitz & formiggini, 2006). the last, 1989 z andtype eruption was described by cassatella et al. (1992), skopal et al. (1997). the uv/optical continuum cooled to ∼ 20 000 k, and its source expanded to ∼ 7 r� having a luminosity of ∼ 10 000 l�. (see also skopal, 2005). the line spectrum showed violet-shifted absorption lines indicating a mass outflow from the hot component at 100–500 km s−1. the recent, 2006 outburst of bf cyg was first reported by munari et al. (2006). spectroscopic observations by sitko et al. (2006), iijima (2006) and mckeever et al. (2011) indicated appearance of strong p-cyg type of hi, hei line profiles from the beginning of the outburst. photometric observations indicated a high level of the star’s brightness for much longer time than was observed for the previous, 1989-92, active phase (siviero et al. 2012; skopal et al., 2012). recently, skopal et al. (2013) reported an evidence of highly-collimated bipolar ejection from bf cyg. 277 http://dx.doi.org/10.14311/app.2015.02.0277 a. skopal et al. 9 10 11 12 13 47000 48000 49000 50000 51000 52000 53000 54000 55000 56000 1990 1995 2000 2005 2010 m ag ni tu de julian date 2 400 000 bf cyg q u i e s c e n t p h a s e a c t i v e p h a s eact. phase v b u figure 1: the ubv light curves of bf cyg from 1986 to the present. they cover the last, 1989-93, and the present, 2006-13, active phases. during the quiescent phase (∼ 1994 − 2006), the star was by 2–3 mag fainter, being characterized with the wave-like orbitally-related variation. in this contribution we analyze our optical spectroscopy and ubv rcic photometry from the current active phase of bf cyg, with the aim to determine fundamental parameters of the hot component. we point the problem of its high luminosity, which sustains for a long time of years. 2 observations broad-band photoelectric ubv and ccd ubv rcic photometry of bf cyg was carried out by 0.6-m telescopes at the skalnaté pleso and stará lesná observatories of astronomical institute of the slovak academy of sciences (see skopal et al. 2012 for details). the data are plotted in fig. 1. the high-resolution spectroscopy was carried out by the single dispersion slit spectrograph mounted at the coudé focus of the 2-m rcc telescope of the rozhen national astronomical observatory and at the ondřejov observatory. the low-resolution (r ∼ 1000) spectroscopic observations were secured by the 2.6-m shajn telescope, operated by the crimean astrophysical observatory. spectroscopic observations were dereddened with eb−v = 0.35 and the resulting parameters were scaled to a distance of 3.8 kpc (e.g. skopal, 2005). 3 analysis and results 3.1 modelling the sed in the optical assuming that the optical continuum consists of the three basic radiative components of radiation (see sect. 1), the resulting flux in the continuum, f(λ), can be expressed as their superposition, f(λ) = fwd(λ) + fn(λ) + fg(λ), (1) where fwd(λ) is the flux from the wd’s pseudophotosphere, fn(λ) is the flux from thermal plasma and fg(λ) represents the flux from the giant. for effective temperatures, t effwd ∼ 5 000 − 10 000 k, an atmospheric model, fλ(t effwd), is needed to fit the radiation of the warm pseudophotosphere. otherwise, a simple blackbody radiation is satisfactory. the nebular radiation in the continuum can be approximated by processes of recombination and thermal bremsstrahlung in the hydrogen/helium plasma for case b. finally, radiation from the giant is represented by an appropriate synthetic spectrum, fλ(t effg ). then eq. (1) can be expressed as, f(λ) = θ2wdfλ(t eff wd) + knελ(te) + θ 2 gfλ(t eff g ), (2) where θwd = rwd/d and θg = rg/d are angular radii of the wd pseudophotosphere and the giant, respectively. the factor kn (= the observed emission measure in cm−5) scales the volume emission coefficient ελ(te) of the nebular continuum to observations. constant electron temperature, te, throughout the nebula is assumed. physical parameters of the model spectrum (2), θwd, θg, t eff wd, t eff g , kn and te, are given by the solution of eq. (2), which corresponds to a minimum of the reduced χ2 function. the sed-fitting analysis was described by skopal (2005) and skopal et al., (2011). 3.2 physical parameters example of a low-resolution (3400 – 7000 å) spectrum, taken around a brightness maximum (23/10/2008), is depicted in fig. 2. the model sed shows that the spectrum is dominated by the radiation of the warm wd pseudophotosphere (denoted as the warm stellar component (wsc) by skopal et al., 2011) and the nebular continuum. the light from the giant becomes more significant for λ > 6 600 å. 278 what powers the 2006 outburst of the symbiotic star bf cygni? the wsc is produced by a source with t effwd ∼ 8 500 k, the effective radius of ∼ 25 r� and the luminosity of ∼ 3000 l�. the nebular component was characterized with a high emission measure of em = 4πd2 × kn ∼ 2.6 × 1061(d/3.8 kpc)2 cm−3, radiated at te ∼ 30 000 k, which correspond to the luminosity ln ∼ 5100 l�. thus the lower limit of the total hot component luminosity was ∼ 8 100 l�, because only a fraction of the burning wd radiation can be converted to the wsc and the nebular emission. 0 1 2 3.6 3.7 3.8 3.9 lo g( f lu x) [ er g cm -2 s -1 a -1 ] + 1 3 log(wavelength) [a] bf cyg23/10/2008 h e i h α h β h γ h δ h e i obs. fg(λ) fwd(λ) fn(λ) sed figure 2: an example of the low-resolution spectrum (gray line) and its model (heavy solid line) taken during the 2006-13 active phase of bf cyg, on 23/10/2008. the model sed and its components of radiation here represent a graphic form of eq. (1) with the same denotation in keys. 3.3 a disk-like shape of the wd pseudophotosphere the shape the wd pseudophotosphere cannot be spherical, because of the simultaneous presence of the strong nebular emission in the spectrum. if it were a sphere, its radiation would not be capable of giving rise to the observed nebular emission. on the other hand, the presence of the strong nebular emission in the spectrum constrains the presence of a hot ionizing source in the system. this type of the spectrum (called as twotemperature type) suggests that the wd pseudophotosphere has a form of a disk. when viewing the disk under a high inclination, its outer rim simulates the warm photosphere (producing the wsc), while the material above/below the disk is ionized by the hot central source and thus converts a fraction of its radiation to the nebular emission (see skopal 2005 and skopal et al. 2011 in detail). 3.4 collimated mass ejection figure 3 shows evolution of the hα profiles from the 2006 august eruption to 2013 april. the broad wings expanding to ∼ ±2000 km s−1 in hα were present in all spectra. significant variations were observed mainly at/around the line cores. (i) during the whole active phase, a sharp absorption developed on the blue side of the profile. (ii) during 2009, additional satellite emission components appeared at the position of a few times ±100 km s−1 to the hα emission core. (iii) during 2012 september, the satellite components were placed nearly symmetrically with respect to the hα central emission. 1 2 3 4 5 6 7 -2000 -1000 0 1000 2000 lo g( f lu x) + c on st . heliocentric radial velocity [km s-1] 30/08/2006 05/06/2009 02/09/2012 bf cyg hα dd/mm/yyyy +5.0 +4.1 +2.5 +1.0 const. 03/11/2012 [n ii ] f e ii 24/04/2013 figure 3: evolution in the hα line profile along the outburst. the filled curves represent the jet emission components (sect. 3.4). fluxes are in 10−13 erg cm−2 s−1 å−1. their radial velocities of ± ∼ 370 km s−1 and fluxes of ∼ 1.4 × 10−11 erg cm−2 s−1 were around a maximum (see skopal et al., 2013 in detail). (iv) the presence of satellite components and their properties were unstable in the spectrum. during two months after their 279 a. skopal et al. best pronounced stage (on 02/09/2012), they practically disappeared on 2012 november 3rd. however, in 2013 april, they re-appeared again (fig. 3). the relatively small width of the (well measured) satellite components (fwhm ∼ 245 km s−1) and their radial velocities suggest that these emissions were produced by radiation of a highly collimated ejection by the central star. 4 concluding remarks according to the elements of the spectoscopic orbit (fekel et al., 2001), the mass of the wd in bf cyg is as low as ∼ 0.55 − 0.6 m�. during the quiescent phase, the luminosity of the hot component was estimated to ∼ 10 000 l� for d = 3.8 kpc (e.g. mikolajewska et al., 1989). this quantity suggests that the source of such the energy output is caused by a stable hydrogen burning on the wd surface at the accretion rate of ∼ 1.4×10−7 m� yr−1 for the 0.55−0.6 m� wd (e.g. shen & bildsten, 2007). during active phase, the luminosity of the hot component can be > 10 000 l� (e.g. cassatella et al., 1992), however, with difficulties of its precise determination as mentioned in sect. 3.2. in addition, (i) a significant extension and thus cooling of the wd pseudophotosphere is indicated by modelling the sed (sect. 3.2), (ii) an enhanced mass-loss rate from the wd is evidenced by the broad hα wings with a violet-shifted absorption component, and (iii) an enhancement of the accretion rate onto the wd is required by the satellite emission components. the presence of bipolar jets confirms the presence of a disk around the accretor during the outburst. these observational properties are consistent with evolution of burning wds in the h-r diagram, when the accretion rate increases above the stable burning regime. the accretion at ≈ 2 × 10−7 m� yr−1 throughout the disk during the outburst can sustain the high luminosity of the burning wd at ≈ 10 000 l� (see fig. 2 of shen & bildsten, 2007). the case of the current bf cyg active phase lead us to a speculation that the simultaneous presence of the enhanced mass outflow and mass infall during some active phases of sss can reflect a new type of the accretion process, which can sustain a high luminosity of their hot components for a long time of years. similar properties with mass outflow/infall and jets were also observed during the 1977–1984 active phase of ch cyg (e.g. skopal et al. 2002). it is of interest to note that the enhanced mass outflow, sometimes followed with jet-like components, and emergence of a warm pseudophotosphere simulated by the irradiated disk are also observed during optical high states of supersoft x-ray sources (e.g. southwell et al. 1996; hutchings et al. 2002; hachisu & kato, 2003). acknowledgement this research was supported by the slovak-bulgarian research and development cooperation project skbg-0015-10, by the grant bstc no. 01/14 bulgariaslovakia, by the grant do 02-85 of bulgarian scientific research fund, by the research program msm0021620860 of the ministry of education of the czech republic and by a grant of the slovak academy of sciences vega no. 2/0002/13. references [1] cassatella a., et al.: 1992, a&a 258, 368 [2] fekel, f. c., et al.: 2001, aj 121, 2219 [3] hachisu, i., kato, m.: 2003, apj, 588, 1003 doi:10.1086/374303 [4] hutchings et al.: 2002, aj, 124, 2833 [5] iijima, t.: 2006, cbet no. 633 [6] leibowitz, e. m., formiggini, l.: 2006, mnras 366, 675 doi:10.1111/j.1365-2966.2005.09895.x doi:10.1086/662076 [7] mckeever, j., et al.: 2011, pasp 123, 1062 [8] mikolajewska, j., et al.: 1989, aj 98, 1427 [9] munari, u., et al.: 2006, cbet no. 596 [10] mürset, u., schmid, h. m.: 1999, a&as 137, 473 doi:10.1086/513457 [11] pucinskas, a.: 1970, bull. vilnius univ. astron. obs. no. 27, 24 [12] shen, k. j., bildsten, l.: 2007, apj 660, 1444 [13] sitko, m. l., et al.: 2006, iauc no. 8746 [14] siviero, a. et al.: 2012, baltic astron. 21, 188 [15] skopal a.: 1992, ibvs no. 3780 doi:10.1093/mnras/292.3.703 [16] skopal, a.: 2005, a&a 440, 995 doi:10.1046/j.1365-8711.2002.05715.x [17] skopal, a., et al.: 1997, mnras 292, 703 [18] skopal, a., et al.: 2002, mnras 335, 1109 doi:10.1002/asna.201111655 [19] skopal, a., et al.: 2011, a&a 536, id. a27 [20] skopal, a., et al.: 2012, astron. nachr. 333, 242 doi:10.1086/177931 280 http://dx.doi.org/10.1086/374303 http://dx.doi.org/10.1111/j.1365-2966.2005.09895.x http://dx.doi.org/10.1086/662076 http://dx.doi.org/10.1086/513457 http://dx.doi.org/10.1093/mnras/292.3.703 http://dx.doi.org/10.1046/j.1365-8711.2002.05715.x http://dx.doi.org/10.1002/asna.201111655 http://dx.doi.org/10.1086/177931 what powers the 2006 outburst of the symbiotic star bf cygni? [21] skopal, a., et al.: 2013, a&a 551, id. l10 [22] southwell et al.: 1996, apj, 470, 1065 discussion dmitry bisikalo: if you have a wind from the wd during the quiescence, how can you accumulate matter to form a huge accretion disk? augustin skopal: the presence of a wind from the wd is indicated even during quiescent phase by, for example, the broad hα wings. the presence of a large disk-like formation during active phases is observationally confirmed. however, its creation is not well understood yet. 281 introduction observations analysis and results modelling the sed in the optical physical parameters a disk-like shape of the wd pseudophotosphere collimated mass ejection concluding remarks 205 acta polytechnica ctu proceedings 2(1): 205–211, 2015 205 doi: 10.14311/app.2015.02.0205 theory of nova outbursts and type ia supernovae m. kato1, i. hachisu2 1department of astronomy, keio university, hiyoshi, yokohama 223-8521, japan 2department of earth science and astronomy, college of arts and sciences, the university of tokyo, komaba, meguroku, tokyo 153-8902, japan corresponding author: mariko@educ.cc.keio.ac.jp abstract we briefly review the current theoretical understanding of the light curves of novae. these curves exhibit a homologous nature, dubbed the universal decline law, and when time-normalized, they almost follow a single curve independently of the white dwarf (wd) mass or chemical composition of the envelope. the optical and near-infrared light curves of novae are reproduced mainly by free-free emission from their optically thick winds. we can estimate the wd mass from multiwavelength observations because the optical, uv, and soft x-ray light curves evolve differently and we can easily resolve the degeneracy of the optical light curves. recurrent novae and classical novae are a testbed of type ia supernova scenarios. in the orbital period versus secondary mass diagram, recurrent novae are located in different regions from classical novae and the positions of recurrent novae are consistent with the single degenerate scenario. keywords: novae light curve analysis wd mass uv x-rays type ia supernova. 1 introduction a nova outburst is a thermonuclear runaway event on a white dwarf (wd). after unstable hydrogen shellburning sets in, the envelope greatly expands and the wd becomes very bright and reaches its optical maximum. after the maximum expansion of the photosphere, the wd’s radius shrinks with time. free-free emission of the expanding ejecta dominates the spectrum in the optical and infrared regions. a large part of the hydrogen-rich envelope is blown away in the wind and the photosphere gradually moves inside. figure 1: light curves of various speed classes of novae. from left to right, v1500 cyg, v838 her, v1668 cyg, pw vul, v723 cas, and pu vul (horizontal line and inset), which are taken from kato et al. (2013). figure 1 shows v and visual light curves of six wellobserved galactic classical novae with different speed classes. v1500 cyg is an exceptionally bright (superbright), very fast nova (t2 = 2.9 days and t3 = 3.6 days). here, t2 (t3) is the time in days in which a nova decays by two (three) magnitudes from its maximum. v838 her is a very fast nova, one of the fastest novae except for the superbright novae. it exhibits a normal super-eddington luminosity in the early phase. v1668 cyg is a fast nova with t2 = 12 days and t3 = 25 days. pw vul is a slow nova that exhibits oscillatory behavior in the early phase. v723 cas is a very slow nova that features multiple peaks in the early stage. pu vul is a symbiotic nova with a flat optical peak lasting eight years. its light curve for the first 20 years is plotted in the inset of figure 1. many classical novae exhibit super-eddington luminosity, i.e., the peak luminosity exceeds the eddington luminosity at the photosphere. ledd ≡ 4πcgmwd κ , (1) where κ is the opal opacity. if we assume mwd = 1.0 m� and κ = 0.4, the eddington luminosity is ledd = 1.258 × 1038 erg sec−1. because of space limitations, in this paper we concentrate on the light curve analysis of the decay phase of novae and their relation to type ia supernova (sn ia) progenitors. for other important issues readers may re205 http://dx.doi.org/10.14311/app.2015.02.0205 m. kato, i. hachisu fer to a brief summary of super-eddington luminosity and a theoretical explanation of maximum magnitude versus rate of decline of novae (the mmrd relation) by kato (2012), the common path in the color-color diagram by hachisu and kato (2014), and information on instability of the shell flash by kato & hachisu (2012). 2 nova light curves 2.1 the universal decline law decay phases of novae can be followed based on the theory of optically thick winds. in this method, we construct a sequence of steady-state solutions with assumed chemical composition of the hydrogen-rich envelope. we solve the equations of motion, mass continuity, radiative diffusion, and conservation of energy, from the bottom of the hydrogen-rich envelope through the photosphere assuming a steady-state (kato & hachisu 1994). optical and infrared light curves are calculated by assuming free-free emission originating from the optically thin ejecta outside the photosphere. the flux of free-free emission is estimated by fν ∝ ∫ nenidv ∝ ∫ ∞ rph ṁ2wind v2windr 4 r2dr ∝ ṁ2wind v2phrph , (2) during the optically thick wind phase, where fν is the flux at frequency ν, ne and ni are the number densities of electrons and ions, respectively, rph is the photospheric radius, ṁwind is the wind mass-loss rate, vph is the photospheric velocity, and ne ∝ ρwind and ni ∝ ρwind. these ṁwind, rph, and vph values are calculated from our optically thick wind solutions. figure 2 shows light curves calculated for various wd masses. as the wind mass-loss rate quickly decreases with time, the flux at a given wavelength also decreases with time. note that the shape of the light curve is independent of the frequency (or wavelength), whereas the absolute magnitude, i.e., proportionality constant of equation (2), depends on it. this wavelength-free light curve shape is one of the characteristic properties of free-free emission (see hachisu & kato 2006). for a more massive wd, its light curve decays more quickly mainly because of its smaller envelope mass, so the evolution time scale is shorter. these light curves follow a common shape, known as the universal decline law. if we plot these light curves on logarithmic time since the outburst, and shift them in the vertical and horizontal directions (i.e., normalize them in the time direction by a factor of fs and also shift the magnitude in the vertical direction by −2.5 log fs), all the light curves essentially converge into a single curve (hachisu & kato 2010) independently of the wd mass and chemical composition, as demonstrated in figure 3. this figure also shows the uv 1455 å narrowband light curve. this narrow band is defined observationally to avoid prominent emission and absorption lines in the nova uv spectra and to represent the continuum flux (cassatella et al. 2002). we see that our model light curves of the uv 1455 å band also converge into a single curve by using the same timescaling factor of fs. in our model light curves, x-ray turnoff times do not converge, because the supersoft xray phase (duration of steady nuclear burning) depends on the hydrogen burning rate, which does not follow the times-scaling law. figure 2: free-free emission model light curves for various wd masses from 0.55 m� to 1.2 m� in steps of 0.05 m� (taken from hachisu & kato 2010). open squares denote observational y magnitudes of v1668 cyg that show good agreement with the 0.95 m� model. this figure also shows the fundamental properties of nova evolution. after the optical maximum, the photospheric temperature rises with time, and the main emitting wavelength region shifts from optical to uv and then to supersoft x-ray. the optically thick winds continue until the photospheric temperature reaches log t (k) > 5.2 (which corresponds to the opacity peak from iron ionization). the wd photosphere emits supersoft x-rays until hydrogen nuclear burning stops. after the hydrogen burning is extinguished, the nova enters a cooling phase and finally the wd becomes dark. figure 4 shows the same light curves as in figure 3 but on a real time scale. in a light curve analysis, we can select the best-fit theoretical model by comparing observed data with our model, but only the optical light curve is not enough. this is because the main body of the optical light curve follows the same decline rate of t−1.75 and we can fit several light curves of different wd masses with the optical data as seen in figure 4. if we have information other than optical, such as the uv 1455 å light curve or x-ray turn-on and turnoff times, 206 theory of nova outbursts and type ia supernovae we can accurately select one model light curve. this is why multiwavelength observations are important for determining the nova parameters. figure 3: free-free emission model light curves for various wd masses from 0.7 m� to 1.3 m� by steps of 0.05 m� (taken from hachisu & kato 2010 ). the time scales are squeezed or stretched by a factor of fs along time to fit each other and the magnitude is shifted by −2.5 log fs. the two straight lines indicate the decline rates of free-free flux, i.e., fν ∝ t−1.75 (dash-dotted line) and fν ∝ t−3 (solid line). observational data of v1974 cyg are superposed. 2.2 chemical composition of ejecta classical novae exhibit heavy element enrichment in their ejecta by a few to several tens of percent by mass (see table 1 in hachisu & kato 2006 for a summary). this enhancement is considered to originate the wd by diffusion of accreted hydrogen into a core during a quiescent phase (e.g., prialnik 1986). as the surface of a wd core is eroded during nova outbursts and blown off in the nova wind, the wd mass decreases after each nova outburst. however, in recurrent novae, the composition of ejecta is very different from those of classical novae and is almost solar. theoretical calculations showed that a certain amount of processed helium is added to the wd after every outburst (prialnik & kovetz 1995). the wd could then experience repeated weak helium shell flashes. the resultant wind is weak, and a large part of the envelope mass remains on the wd (kato & hachisu 2004). therefore, the wd increases its mass in recurrent novae such as u sco and rs oph. note that helium burning produces a substantial amount of ne and mg (shara & prialnik 1994). therefore, even if the wd is made of carbon and oxygen, the hydrogen-rich envelope could be contaminated with neon. mason (2011) claimed that the high ne/o line ratios observed in u sco indicate that u sco harbors an onemg wd. mikolajewska criticized mason’s results because mason neglected the effects of collisions on the line formation. with these effects included, the u sco abundance is consistent with the solar abundance ratio. mason (2013) accepted mikolajewska’s criticism and revised her results on the u sco wd, which previously had been totally erroneous because of her misinterpretation and miscalculation. here, we note that (1) high ne/o line ratios could be an indication of a massincreasing wd and (2) strong ne lines are detected even in the slow nova v723 cas (iijima 2006), which harbors a low-mass wd (0.5–0.6 m�). figure 4: same as figure 3 but on a real time scale and using absolute v magnitude (taken from hachisu & kato 2010). model uv and supersoft x-ray light curves are plotted only for three wd masses (1.1 m�, 1.05 m�, and 1.0 m�) to improve readability. 2.3 wd masses estimated by light curve fittings figures 3 and 4 also show light curves of the classical nova v1974 cyg. the optical light curve decays as t−1.75 from the optical peak at t = 2.67 days (point b). there are two different sets of v -magnitudes obtained by using different v filters that have different degrees of emission line contamination. as the nova enters the nebular phase, strong emission lines such as [o iii] contribute to the v band flux and the v magnitude becomes much brighter than the continuum flux. (the theoretical free-free flux mimics the continuum flux.) after the optically-thick winds stops, the magnitude decays as t−3, which represents free-expansion of the ejecta. because the optical, uv, and supersoft x-ray light curves have different dependencies on the wd mass, we can choose a best-fit model that reproduces these three light curves simultaneously as shown in figure 4. in this way, we have estimated the wd mass in many classical novae, recurrent novae, and a symbiotic nova. such 207 m. kato, i. hachisu examples are shown in figure 1 by the solid lines that represent theoretical light curve models. these wd masses are mwd = 1.20 m� for v1500 cyg, 1.35 m� for v838 her, 0.95 m� for v1668 cyg, and 0.6 m� for v723 cas. note that some classical novae have optical light curves that are similar to those of recurrent novae. for example, v838 her and u sco have very similar optical and uv 1455 å light curves. also the classical nova v2491 cyg exhibits a similar optical decline to that of rs oph. v838 her and v2491 cyg are classical novae and their wd masses are decreasing because of heavy element enhancement in their ejecta, whereas u sco and rs oph have mass-increasing wds. this indicates that classical and recurrent novae share different histories of binary evolution even if they exhibit very similar light curves. 2.4 very slow novae with a flat optical peak pu vul is a symbiotic nova that exhibits a long-lasting flat optical peak followed by a slow decline, as shown in figure 1. the evolution of this nova can be followed by using a quasi-evolution model of an outburst on a ∼ 0.6 m� wd (kato et al. 2011). in less massive wds (≤ 0.6 m�, which depends slightly on the chemical composition) the optically thick winds are not accelerated and thus the nova evolution is extremely slow. in contrast, in the majority of classical novae (that occur on a wd of mwd > 0.7 m�) the strong optically thick winds are accelerated and carry most of the envelope mass in a short time. as a result, their evolutions are fast and exhibit a sharp optical maximum. the very slow novae v723 cas, hr del, and v5558 sgr exhibit a flat pre-maximum phase with no indication of a strong wind-mass loss, followed by a smooth decline with massive winds after some oscillatory phase of the optical light curves. this sequence can be understood as a transition from initial quasi-static evolution to optically thick wind evolution. the oscillatory behaviors in the light curves may correspond to a relaxation process associated with the transition from static to wind solutions (kato and hachisu 2011). 3 type ia supernova scenarios type ia supernovae are characterized in principle by spectra without hydrogen lines but with strong si ii at maximum light. it is commonly agreed that the exploding star is a mass-accreting carbon-oxygen (c+o) wd. however, whether the wd accretes hand he-rich matter from its binary companion (the so-called single degenerate (sd) scenario) or whether two c+o wds merge (the so-called double degenerate (dd) scenario) has not yet been clarified. in the dd scenario, proposed by webbink (1984) figure 5: initial and final binary parameter ranges for the sd scenario (taken from hachisu et al. 2012). m2 is the companion mass and porb is the orbital period. an initial binary system inside the region encircled by a solid line (labeled initial) evolves to explode as a type ia supernova when it reaches within the region labeled final. the recurrent novae u sco and t crb are located inside the final region, whereas the classical novae v2491 cyg, v838 her, and gk per lie outside. most classical novae and cataclysmic variables are located much below the left-side final region. 208 theory of nova outbursts and type ia supernovae and iben & tutukov (1984), intermediate-mass (3– 8 m�) binaries undergo the first and second common envelope phases and finally become a double wd system. if the orbital period is very short (i.e., if they merge within a hubble time), and the total mass exceeds the chandrasekhar mass, it could explode as a sn ia. after introduction of this traditional dd scenario, many modified binary evolution scenarios have been proposed, but the evolutionary path is essentially governed by the efficiency parameter of the common envelope evolution (e.g. webbink 2008), which is not well constrained yet. the original idea of the sd scenario was proposed by whelan & iben (1973), but this idea was soon abandoned by one of its originators (iben) in 1984 because of its too small contribution to sne ia. these studies are based on the old opacity, which was revised at the beginning of the 1990s. the new opacities have a large peak at log t (k) ∼ 5.2, which brought drastic change in stellar structures. the presence of strong optically thick winds has changed nova theory, and we now can well reproduce observational light curves. hachisu et al. (1996) adopted these optically-thick winds as a new elementary process in binary evolution, which was dubbed the accretion wind evolution. the wd accretes matter from the accretion disk and, at the same time, blows optically thick winds. these winds carry mass and angular momentum, so mass transfer in the binary can be stabilized, and, in most cases, the so-called second common envelope evolution does not occur. wds can grow in mass to the chandrasekhar mass and explode as a sn ia. therefore, the essential difference between the dd and sd scenarios resides in the assumption of the second common envelope evolution or accretion wind evolution. the physical process of the latter is well studied in nova outbursts, whereas the former process is not established yet (see e.g., webbink 2008, kato and hachisu 2012 for more details). 3.1 wd+rg channel and wd+ms/sg channel in the sd scenario, there are two known channels to sne ia, i.e., the wd+ms (including sub-giant (sg) companion) channel and the wd+rg channel. the wd+rg channel was proposed by hachisu et al. (1996), who called it the symbiotic channel. in this symbiotic channel we start from a very wide binary consisting of a pair of ms stars. the primary evolves to an asymptotic giant branch (agb) star. the binaries undergo a common envelope-like evolution during the superwind phase of the primary agb star. the orbital period of the binary shrinks to 30 – 800 days, with the binary now consisting of a co wd and a ms star. then the secondary ms star evolves to a rg with a helium core and fills its roche lobe. as the mass-transfer rate exceeds the critical rate, the binary undergoes the accretion wind evolution phase. the mass-transfer rate gradually decreases and the optically thick winds stop. then all the accreted matter is burned on the primary co wd. the binary enters a persistent supersoft x-ray source (sss) phase. the mass-transfer rate further decreases to below the minimum rate of steady hydrogen burning (ṁst) and the binary enters a recurrent nova phase. depending on the binary parameters, the wd mass reaches mwd = 1.38 m� during one of the three phases (wind, sss, and recurrent nova) and explodes as a type ia supernova. the ms/sg channel was proposed by li & van den heuvel (1997) and further investigated by hachisu et al. (1999b) and others. in this channel, a pair of ms stars evolves to a binary consisting of a helium star plus a ms star after the first common envelope evolution. the primary helium star further evolves to a helium rg with a co core and fills its roche lobe followed by a stable mass-transfer from the primary helium rg to the secondary ms star. the secondary ms star increases its mass and is contaminated by the primary’s nuclear burning products. the primary becomes a co wd after all the helium envelope is transferred to the secondary ms star. then, the secondary evolves to fill its roche lobe and mass transfer begins from the secondary ms (or subgiant) star to the primary co wd. then, the system enters an accretion wind evolution phase. after that, the binary becomes a persistent sss and evolves finally to a recurrent nova. during these three periods, the primary wd increases its mass to the chandrasekhar mass. depending on the binary parameters, the wd will explode as a sn ia during one of the three phases (wind, sss, and recurrent nova). some binaries are identified by the last phase of each channel. in the symbiotic channel, the smc symbiotic x-ray source smc3 corresponds to the sss phase and symbiotic recurrent novae such as rs oph and t crb correspond to the last phase. in the ms/sg channel, v sge and rx j0513.9−6951 correspond to the accretion wind evolution phase. steady-burning supersoft sources such as cal 87 and cal 83 correspond to the sss phase. u sco and v394 cra are recurrent novae in the ms/sg channel. 3.2 evolutionary status of novae and recurrent novae figure 5 shows the orbital period versus secondary mass diagram for sn ia progenitor binaries for both the wd+ms/sg channel and the wd+rg channel. the upper region labeled “initial” for each channel indicates that binaries will evolve downward with the secondary mass decreasing and the wd mass increasing and even209 m. kato, i. hachisu tually explode as a sn ia in the “final” region. the present positions of u sco and t crb in this diagram are very consistent with the final stage. from light curve analysis, the wd mass is determined as massive as 1.37 m� for u sco and 1.35 m� for rs oph and t crb. these values are also consistent with this final stage toward a type ia sn. in contrast, classical novae v838 her and v2491 cyg are located much below the final region, indicating that neither is on the way toward a sn ia explosion in the wd+ms/sg channel nor in the wd+rg channel, even though they harbor a massive wd near the chandrasekhar mass. in this way, the present status of these recurrent novae and classical novae can be understood from the difference in binary evolution that is consistent with the sd scenario. however, in the original dd scenario, there are no known paths to recurrent novae, especially for the rs oph type systems. thus, studies of cvs are important in understanding the nature of progenitors of sne ia. references [1] cassatella, a., altamore, a., gonzález-riestra, r.: 2002, a&a 384, 1023 [2] della valle, m.: 1991, a&a 252, 9 [3] hachisu, i., kato, m.: 2001, apj 558, 323 doi:10.1086/321601 [4] hachisu, i., kato, m.: 2006, apjs 167, 59 doi:10.1086/508063 [5] hachisu, i., kato, m.: 2010, apj 709, 680 doi:10.1088/0004-637x/709/2/680 [6] hachisu, i., kato, m.: 2014, apj 785, 97 doi:10.1088/0004-637x/785/2/97 [7] hachisu, i., kato, m., nomoto, k.: 1996, apj 470, l97 [8] hachisu, i., kato, m., nomoto, k.: 1999a, apj 522, 487 [9] hachisu, i., kato, m., nomoto, k.: 2008, apjl 683, 127 doi:10.1086/591646 [10] hachisu, i., kato, m., nomoto, k.: 2012, apjl 756, l4 doi:10.1088/2041-8205/756/1/l4 [11] hachisu, i., kato, m., nomoto, k., umeda, h.: 1999b, apj 519, 314 doi:10.1086/307370 [12] iben, i., jr., tutukov, a. v.: 1996, apjs 105, 145 [13] iijima, t.: 2006, a&a 451, 563 [14] kato, m., hachisu, i.: 1994, apj 437, 802 [15] kato, m., hachisu, i.: 2004, apj 613, l129 doi:10.1086/425249 [16] kato, m., hachisu, i.: 2009, apj 699, 1293 doi:10.1088/0004-637x/699/2/1293 [17] kato, m., hachisu, i.: 2011, apj 743, 157 [18] kato, m, hachisu, i.: 2012, basi 40, 393. [19] kato, m., hachisu, i., cassatella, a., gonzález-riestra r.: 2011, apj 727, 72 [20] kato, m, hachisu, i., henze, m.: 2013, apj, 779, 19 [21] mason, e.: 2011, a&a 352, l11 (2013, a&a 556, c2: corrigendum) [22] prialnik, d.: 1986, apj 310, 222 [23] prialnik, d., kovetz, a.: 1995, apj 445, 789 [24] shara, m., m., prialnik, d.: 1994, aj 107, 1542 [25] webbink, r. f.: 1984, apj 277, 355 [26] whelan, j., iben, i., jr.: 1973, apj 186, 1007 [27] webbink, r. f.: 2008, assl 352, 233 discussion jan-uwe ness: for a sn ia, the wd has to be a co wd, thus, the progenitor main-sequence was low mass star. are there really enough co wds with highenough initial mass? furthermore, the companion will initially be of rather low mass in order not to evolve faster than the primary star. can such a companion feed a lot of mass onto the co wd until the sn ia explosion? thus a sd progenitor will have a very low mass companion just before sn ia explosion. mariko kato: a co wd originates from less massive stars (m at zams < 8 m�). the companion mass is smaller than this value. we have calculated the binary evolution including mass exchange, mass-loss from the binary, wind mass-loss from the wd, and so on. as you see in figure 5, the mass permitted for the secondary is rather limited. especially in the wd+rg channel, the initial companion mass is small, i.e., 1 – 3 m� when the secondary had just filled its roche lobe, and 0.5 – 1.0 m� just before the sn ia explosion. for the primary wd, the initial 0.9 – 1.07 m� wd becomes a type ia sn. we calculated the expected number of sne ia and showed that the two channels have enough sne ia. their delay time distribution is also consistent with observation (hachisu et al. 2008). 210 http://dx.doi.org/10.1086/321601 http://dx.doi.org/10.1086/508063 http://dx.doi.org/10.1088/0004-637x/709/2/680 http://dx.doi.org/10.1088/0004-637x/785/2/97 http://dx.doi.org/10.1086/591646 http://dx.doi.org/10.1088/2041-8205/756/1/l4 http://dx.doi.org/10.1086/307370 http://dx.doi.org/10.1086/425249 http://dx.doi.org/10.1088/0004-637x/699/2/1293 theory of nova outbursts and type ia supernovae giora shaviv: we do a calculation of one flash and find that 10−7m� is accreted. here we infer that this system will be a sn ia. but we saw only how a 10−7m� are accreted and infer that the same is true with 0.1 m�. this is an extrapolation over 6 orders of magnitude!! hence can easily be wrong. mariko kato: in our estimate based on the light curve analysis, the accumulation efficiency in typical recurrent novae (u sco and rs oph) is about 50%. the accreted matter in one nova outburst is 1.2 × 10−6m� for u sco, and 1−2×10−6m� for rs oph. as the wd is estimated to be 1.35 m� for rs oph and 1.37 m� for u sco, we can expect a continuous mass increase of the wds, not just from an extrapolation from 1.3 m�. 211 introduction nova light curves the universal decline law chemical composition of ejecta wd masses estimated by light curve fittings very slow novae with a flat optical peak type ia supernova scenarios wd+rg channel and wd+ms/sg channel evolutionary status of novae and recurrent novae acta polytechnica ctu proceedings doi:10.14311/app.2015.1.0057 acta polytechnica ctu proceedings 2:57–61, 2015 © czech technical university in prague, 2015 available online at http://ojs.cvut.cz/ojs/index.php/app on sampling based methods for the dubins traveling salesman problem with neighborhoods petr váňa∗, jan faigl dept. of computer science, czech technical university in prague, technická 2, 166 27 prague, czech republic ∗ corresponding author: vanapet1@fel.cvut.cz abstract. in this paper, we address the problem of path planning to visit a set of regions by dubins vehicle, which is also known as the dubins traveling salesman problem neighborhoods (dtspn). we propose a modification of the existing sampling-based approach to determine increasing number of samples per goal region and thus improve the solution quality if a more computational time is available. the proposed modification of the sampling-based algorithm has been compared with performance of existing approaches for the dtspn and results of the quality of the found solutions and the required computational time are presented in the paper. keywords: dubins vehicle, dubins maneuver, dtsp, dtspn. 1. introduction path planning for a non-holonomic vehicle is a fundamental problem of surveillance mission where an unmanned aerial vehicle (such as a fixed-wing) is requested to visit a given set of locations. the basic model of such a vehicle with the limited turning radius is called the dubins vehicle [1] for which the combinatorial problem of finding optimal sequence of visits to the locations is known as the dubins traveling salesman problem (dtsp) [2]. in this paper, we consider a generalized problem of the dtsp where the particular waypoints to be visited can be selected from a set of possible locations. due to the similarity of this problem with the so-called traveling salesman problem with neighborhoods [3], the problem is called as the dubins traveling salesman problem with neighborhoods (dtspn) [4]. the dtspn is a suitable problem formulation to address surveillance missions with unmanned aerial vehicles, where it is required to take a camera snapshot (or other type of measurement) of each target location. such a snapshot can be acquired from some distance of the target location and thus it is not necessary to visit the location exactly. it is rather a more suitable to select the waypoints in such a way that all locations are covered while the total cost of the final path is minimal. there are several approaches to address the dtsp and also the dtspn in the literature [2, 4–6] including our work [7]. therefore, in this paper, we provide an overview of the existing approaches to address the dtspn and compare their performance according to the trade-off between the solution quality and computational requirements. in particular, we focus on the sampling-based algorithm [4] which is able to provide high quality solutions for very high number of samples. however, more samples increase the computational burden, and therefore, the algorithm is not directly suitable for real-time applications. on the other hand, we can get a quick estimation of the solution quality using only few samples. this has motivated us to modify the original algorithm [4] to provide a first solution quickly that is then improved if a computational time is left. the paper is organized as follows. a brief introduction to the problem background is presented in the next section. the addressed problem is formally introduced in section 3. a brief overview of the existing approaches to solve the dtspn is provided in section 4. the proposed modification of the samplingbased approach is presented in section 5 together with the analysis of its computational complexity. evaluation results of the comparison of the existing approaches with the proposed modification are reported in section 6. finally concluding remarks are denoted to section 7. 2. related work the problem of curvature-constrained path planning has been studied years ago and the fundamental work has been published in 1957 by dubins. in [1], he proved that the optimal path between two configurations q1,q2 ∈ se(2) consists of only straight line segments and segments with the minimum turning radius. he also showed that only 6 maneuvers can be optimal, which can be further divided into two main types: • ccc type: lrl, rlr; • csc type: lsl, lsr, rsl, rsr; where c stands for a circle turn (r – right, l – left) and s for a straight segment. even though the optimal path for dubins vehicle between two configurations is known and it is straightforward to compute, this is not sufficient to directly solve the dtsp. it is because the orientation of the vehicle at the waypoints is not known and thus it must be determined together with the optimal sequence of 57 http://dx.doi.org/10.14311/app.2015.1.0057 http://ojs.cvut.cz/ojs/index.php/app petr váňa, jan faigl acta polytechnica ctu proceedings visits to the waypoints. moreover, the optimal orientation depends on the sequence and vice-versa, which make the problem difficult to address. probably the simplest approach to address the dtsp is the alternating algorithm (aa) proposed in [2]. in this approach, headings are established in the way that even edges are connected by straight line segments and odd edges are connected by dubins maneuvers. in addition, the authors show that the length of the optimal solution of the dtsp can be bounded by ltspκdn/2eπρ, where ρ is minimum turning radius, ltsp is the length of the optimal solution of the euclidean tsp, n is the number of the goals, and κ < 2.658. based on the similar idea, authors of [5] proposed a receding horizon algorithm called the look-ahead (la) approach. the heading is determined with respect to the next point in the sequence. this algorithm investigates each point only once, and therefore, the la approach is very fast and suitable for real-time application while the authors reported better results than the aa. the optimal solution of the dubins planning to visit a given sequence of waypoints that are at the distance longer than 4ρ is presented in [8]. the approach is based on the convex optimization; however, the optimization needs to be solved several times because of possible alternation of the maneuvers directions. the authors bound the number of possible combinations to 2n−2 for n waypoints. the dtspn is a generalization of the dtsp, where each goal is extended by a neighborhood (goal region). as the dtsp is known to be np-hard [6], also its generalization the dtspn is np-hard. 3. problem definition the addressed problem is motivated by surveillance missions with fixed-wing aerial vehicles, which are nonholonomic vehicles due to their kinodynamic constraints. these vehicles are often modeled as the dubins vehicle [1], which can go only forward at a constant speed and has a minimum turning radius ρ. the configuration space of such a vehicle can be represented as se(2), where each configuration q ∈ se(2) is a triplet (x,y,θ), where (x,y) is the vehicle position in a plane and θ ∈ s1 is the orientation of the vehicle. the mathematical model of the dubins vehicle can be formulated as follows:  ẋẏ θ̇   = v   cos θsin θ u ρ   , |u| ≤ 1, (1) where v is the vehicle forward velocity, ρ is the minimum turning radius, and u is the control input, u > 0. now, we can introduce the dtspn [9] a more formally. let r = {r1, . . . rn} be a set of n regions ri ⊂ r2 that are requested to be visited by the dubins vehicle and let σ = (σ1, . . . ,σn) be an ordered permutation of {1, . . . ,n}. we define a projection from se(2) to r2, i.e., p(q) = (x,y), and let qi be an element of se(2) whose projection lies in ri. the dtspn is path planning problem where the dubins vehicle has to visits each region ri while satisfying the kinodynamic constraints of (1). every optimal path has to intersect each goal region ri in at least one configuration qi. hence, we can treat the dtspn as an optimization problem over all possible permutations σ and configurations (q1, . . . ,qn) as follows: problem 3.1 (dtspn). minimize σ,q l(qσn,qσ1 ) + n−1∑ i=1 l(qσi,qσi+1 ) subject to p(qi) ∈ ri, i = 1, . . . ,n where l(qi,qj) is the dubins distance between qi and qj. in this paper, we are focused on the problem with regions ri that are mutually exclusive. hence, we can define the minimum distance constraint dk such that for all i,j ∈{1, 2, . . . ,n}, i 6= j,∀pi ∈ ri,∀pj ∈ rj: ||pi −pj|| > kρ. (2) in particular, we are focused on the problem for which the minimum distance dk is equal to 4ρ, which is denoted as the d4 constraint in the rest of this paper. problem instances of the dtspn with the d4 constraint have special properties that are shown and used in [7] to find high quality solutions. in this paper, we also consider d4 problems and compare the existing solutions with the sampling based methods. an overview of the existing methods is provided in the next section. 4. existing methods in literature, we can find several different approaches to address the dtspn. the existing approaches can be divided into three main classes: 1) the decoupled methods; 2) sampling-based methods; 3) and evolutionary methods. representatives of each particular class are briefly described in the following subsections. 4.1. decoupled methods a decoupled method addresses the dtspn by decomposition of the problem into two sub-problems. first, the permutation of the visits to the requested areas (locations) is found independently on the dubins path planning. the second sub-problem is to find a particular visiting configuration of each location. in this way, the original dtspn is transformed to the dtsp with point locations to visit. in [10], the authors address the dtspn by solving the euclidean tsp where cities represents centers of the regions to be visited. once such a permutation is found, the final solution is constructed using the aa [2]. 58 vol. 2/2015 the dubins traveling salesman problem with neighborhoods authors of [7] also consider a solution the etsp to estimate the permutation of the visits. but instead of simple aa, the proposed local iterative optimization (lio) is used to find a significantly better solution. 4.2. sampling-based methods sampling-based methods [4, 11] are another class of existing approaches to address the dtspn. similarly to the decoupled methods, also sampling-based methods are based on existing approaches for the tsp variants; however, in this case a single representative of each region is not considered as in [10] but each region is rather sampled by a particular number of random samples (including the orientation of the vehicle). the samples can be determined in the whole region or only on its boundary. then, the samples of different regions are connected by the dubins maneuver to form a roadmap. such a roadmap is considered as an instance of the generalized asymmetric tsp (gatsp). if the gatsp is solved to the optimum, the sampling-based methods are resolution complete [11]. since the optimal solution of the gatsp can be computationally very demanding, existing heuristic for the tsp can be a more suitable option. the generated gatsp can be solved by its transformation to the asymmetric tsp (atsp) using the noon-bean transformation [12]. then, the transformed atsp can be solved by existing algorithms, e.g., the linkernighan heuristic algorithm [13]. 4.3. evolutionary methods the last class of existing approaches are genetic programming based algorithms. the general idea of these methods is to encode a solution by a chromosome in which the goal permutation and the configurations of the visits are stored. similarly to another approaches, only configurations on the boundary of the regions are considered in these evolutionary methods. probably the first evolutionary approach to the dtspn was published in [14]. the authors adapted the ordered crossover operator (ox) and added two new mutation operators (orientation shift and position shift) to optimize visiting configurations. relatively recently, the genetic approach from [14] was modified into a memetic algorithm in [15]. the authors used similar operators and added a local improvement of individuals in the population by optimizing the position of visiting points. the authors used aa as heuristic algorithm to speed up the algorithm. 5. modification of existing sampling-based algorithm the main issue of sampling based algorithms for the dtspn is that they require a fixed number of samples, which need to be established in advance. we propose an iterative schema to avoid this issue and set the number of samples progressively, which provides first solutions very quickly. moreover, if there is some additional computational time, the algorithm can iteratively improve the solution by adding more samples, which results to obtain a first solution quickly that is further improved. in the basic sampling-based algorithm, a roadmap with (n·m)2 edges is created where n is the number of regions to be visited and m is the number of samples per region. for each edge one dubins maneuver is constructed. since the noon-been transformation does not change the number of vertices, the overall time complexity of generating an instance of the atsp can be bounded by o((n ·m)2). a solution of the generated atsp can be found by the available lkh solver. according to the author of the lkh, the time complexity of the lkh is approximately o(n2.2atsp ) [16]. hence, the expected total time complexity is o((n ·m)2.2). the time complexity to solve the atsp is greater than a roadmap generation, and therefore, we do not need to explicitly consider the construction of the roadmap in the estimation of the required computational time based on the number of samples. based on the relation of the required number of operations on the number of samples, we can establish a sequence of the increasing numbers of the samples needed to find an initial solution quickly and improve its quality later. the sampling-based algorithm works with the constant number of samples given a priory, we can run the sampling based algorithm repeatedly with an increasing number of samples. since the time complexity of the algorithm is nearly quadratic, an inverse function for the number of samples mk according to the number k of the particular iteration can be defined as: mk ≈ √ 2k. (3) after rounding the mk to an integer value, it gives us the following series of the numbers: m = {1, 2, 3, 4, 6, 8, 11, 16, 23, 32, 45, 64, . . .}. (4) figure 1 provides influence of the required computanumber of samples per region 0 100 200 300 400 500 600 700 800 t im e [ s] 10 -2 10 -1 10 0 10 1 10 2 10 3 etsp + lio sample + atsp figure 1. the average required computational time (from 20 trials) for the instance of the dtspn with 4 circle regions and d4 constraint. tional time on the number of samples per each region 59 petr váňa, jan faigl acta polytechnica ctu proceedings to visit. we plot the required computational time on the number of samples per each region to visit. in this case, a simple problem with 4 regions and d4 constraint is considered and the computational time is compared with the etsp+lio algorithm proposed in [7]. 6. results the performance of the proposed modification of the sampling based algorithm has been evaluated in a series of scenarios. the evaluated instances of the dtspn were generated randomly for convex regions satisfying the d4 constraint. several shapes of the regions have been considered: points, disks with the radius equal to ρ, ellipses with the semi-axis 2ρ and 0.5ρ, and convex polygons with up to 6 vertices created from a circle with the radius ρ. a relatively high and uniform density of the regions for the d4 constraint has been considered by generating centers of the regions inside a bounding box with the side 6 √ nρ, where n is the number of the regions to be visited by the dubins vehicles. an example of the examined problems is depicted in figure 2. (a) . d4 convex regions (b) . d1 convex regions figure 2. examples of the randomly generated instances of the dtspn. the examined algorithms are denoted as follows. the newly proposed modification of the samplingbased algorithm is denoted sample+atsp. the algorithm is compared it with the evolutionary approaches [14] and [15] denoted as genetic and memetic, respectively. in addition, we implemented three variants of the decoupled approach based on the solution of the etsp utilized as the heuristic estimation of the permutation of the visits to the regions. the first decoupled approach is denoted as etspn+aa and it is based on the alternating algorithm (aa) [2]. the second method is denoted as the etsp+lio [7] and is based on local optimization of position and heading of the waypoints. the last method is derived from the etsp+lio, but only local iterative optimization of the heading is considered, this method is denoted as the etspn+holio. the quality of found solutions has been evaluated regarding a dedicated computational time for problems with n = 20 and n = 40 regions with the d4 constraint. the quality is measured as the ratio of the path length found by the particular algorithm to the solution provided by the etsp+lio algorithm, similarly as in [7]. all instances were generated randomly, and therefore, 50 trials have been solved for each problem and the algorithm variants. the achieved results are depicted in figure 3 for the 20 regions and in figure 4 for the 40 goal regions. time [s] 10 -2 10 -1 10 0 10 1 10 2 r e la ti v e l e n g th o f th e s o lu ti o n 1 1.1 1.2 1.3 1.4 1.5 etsp + lio etspn + holio etspn + aa memetic genetic sample + atsp figure 3. average ratio of the tour length (from 50 trials) according to the etsp+lio solution for the dtspn with n=20 convex regions. plots start from the time when the first solution is available. time [s] 10 -2 10 -1 10 0 10 1 10 2 r e la ti v e l e n g th o f th e s o lu ti o n 1 1.1 1.2 1.3 1.4 1.5 etsp + lio etspn + holio etspn + aa memetic genetic sample + atsp figure 4. average ratio of the tour length (from 50 trials) according to the etsp+lio solution for the dtspn with n=40 convex regions. plots start from the time when the first solution is available. all the algorithms have been implemented in c++ and tested on a single core of the intel core i5-m480 cpu running at 2.67 ghz. the processor was accompanied with 4 gb ram. 6.1. discussion the presented results show that the proposed modification of the sampling-based algorithm can be considered as a meaningful alternative to the evolutionary based algorithms. in the case of the high number of regions to visit (n = 40), the modified samplingbased algorithm has even superior results to the both genetic and memetic algorithms, while it is less computationally demanding. in the results depicted in figures 3 and 4, the etsp+lio algorithm achieved the best results among 60 vol. 2/2015 the dubins traveling salesman problem with neighborhoods the evaluated algorithms. this is caused by the fact, that all instances of the dtspn were generated with non-overlapping regions with the d4 constraint and the etsp+lio has been designed on top of the properties of the optimal solution of such a problem and thus it directly searches for good solutions [7]. 7. conclusions in this paper, we proposed a modification of the sampling-based algorithm for the dtspn. the proposed algorithm repeatably executes the original sampling-based algorithm with an increasing number of the samples per each region. by this modification, the newly developed algorithm provides the first solution very quickly that can be further improve if more time is available. this makes the algorithm suitable in situations where a solution has to be found quickly and where the maximum time that can be dedicated for the computation is not known in advance. we also compared the modified algorithm with other existing approaches on the randomly generated instances of the dtspn with the d4 constraint. the modified algorithm provides better results than the implemented evolutionary algorithms while it is less computationally demanding. a further comparison of the algorithms’ performance in the instances of the dtspn where regions to visit are closer or overlapping is a subject of our future work. acknowledgements the presented work has been supported by the czech science foundation (gačr) under research project no. gp13-18316p. references [1] l. e. dubins. on curves of minimal length with a constraint on average curvature, and with prescribed initial and terminal positions and tangents. american journal of mathematics pp. 497–516, 1957. doi:10.2307/2372560. 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[cited 31 mar 2015]. 61 http://dx.doi.org/10.2307/2372560 http://dx.doi.org/10.1109/acc.2005.1470055 http://dx.doi.org/10.1109/acc.2011.5991501 http://dx.doi.org/10.1109/cdc.2006.376928 http://dx.doi.org/10.1109/tac.2011.2166311 http://dx.doi.org/10.1137/100816079 http://dx.doi.org/10.3390/a6010084 http://dx.doi.org/10.1109/iecon.2012.6389140 http://dx.doi.org/10.2514/1.48949 http://dx.doi.org/10.1287/opre.39.4.623 http://dx.doi.org/10.1287/opre.21.2.498 http://dx.doi.org/10.2514/6.2009-5888 http://dx.doi.org/10.1016/j.cja.2014.04.024 http://www.akira.ruc.dk/~keld/research/lkh acta polytechnica ctu proceedings 2:57–61, 2015 1 introduction 2 related work 3 problem definition 4 existing methods 4.1 decoupled methods 4.2 sampling-based methods 4.3 evolutionary methods 5 modification of existing sampling-based algorithm 6 results 6.1 discussion 7 conclusions acknowledgements references 163 acta polytechnica ctu proceedings 1(1): 163–169, 2014 163 doi: 10.14311/app.2014.01.0163 modelling the multifrequency sed of agn candidates among the unidentified egret and fermi gamma-ray sources pieter j. meintjes1, pheneas nkundabakura2,1, brian van soelen1, alida odendaal1 1department of physics, university of the free state, p.o. box 339, bloemfontein, 9300, south-africa 2kigali institute of education, p.o. box 5039, kigali, rwanda corresponding author: meintjpj@ufs.ac.za abstract of the 271 sources in the 3rd egret catalogue, 131 were reported as unidentified, i.e. not associated with any particular class of point source in the sky. since the largest fraction of the egret sources were extragalactic, a sample of 13 extragalactic unidentified sources have been selected for multi-wavelength follow-up studies. five of the selected egret sources coincide with gamma-ray flux enhancements seen in the fermi-lat data after one year of operation. in this article, we report the multi-wavelength properties of, among others, the 5 sources detected by fermi-lat from our sample of high galactic latitude unidentified egret sources. recent spectroscopic observations with the southern african large telescope (salt) confirmed one of the unidentified egret sources as a possible seyfert 2 galaxy, or alternatively, a narrow line radio galaxy. the detected gamma-ray emission (eγ > 30 mev) of the 5 coinciding egret/fermi-lat sources are fitted with external compton and synchrotron self compton (ssc) models to investigate the energetics required to produce the egret/fermi gamma-ray flux. in all the models the inclination angle of the jet with respect to the observer is θjet ≈ 60◦, between those of seyfert 1 and seyfert 2/radio galaxies. these results confirm the possibility of seyfert and radio galaxies sources are constituting a new class of γ-ray source in the energy range eγ > 30 mev. keywords: radiation mechanisms: non-thermal line: identification techniques: spectroscopic galaxies: jets bl lacertae objects. 1 introduction the energetic gamma ray telescope experiment egret (30 mev 10 gev) provided the highest gamma-ray window on board the compton gammaray observatory (cgro). egret detected 271 gamma-ray sources above 100 mev, 92 % of which were blazars. of the 271 sources detected, 131 remained unidentified, i.e. could not be associated with any specific point source of gamma-ray emission (hartman et al., 1999). the large number of unidentified egret sources above and below the galactic plane inspired a search for possible extra-galactic radio loud active galactic nuclei (agn) counterparts that could possibly be associated with these unidentified sources. to avoid confusion with possible galactic sources, especially molecular cloud distributions, the search for counterparts was restricted to those unidentified sources at galactic latitudes | b | > 10◦. source selection criteria: the candidate counterparts should be inside the error box associated with the egret detection (hartman et al., 1999), confirmed as extragalactic in the nasa extragalactic database (ned), possess a radio brightness above 100 mjy at 8.4 ghz (sowards-emmerd, romani & michelson, 2003), exhibit hard spectra with spectral indices | α | < 0.7 (sowards-emmerd, romani & michelson, 2003) and display variability (e.g. fan, 2005) that may be associated with an inner accretion disc or jet. based upon these criteria, 13 sources have been selected for further follow-up study (meintjes & nkundababura, 2012; nkundabakura & meintjes, 2012). of the 13 candidate sources selected, 5 have confirmed gamma-ray excesses in the fermi-lat catalogue containing the first year’s observations (abdo et al., 2010a). the fermi large area telescope (fermilat) is a pair conversion γ-ray telescope sensitive to photon energies between 20 mev and 300 gev. launched on 11 june 2008, fermi-lat started to collect data in august 2008. the data are made available on a daily basis and can be accessed online at the official website of the fermi science support centre (http://fermi.gsfc.nasa.gov/cgi-bin/ssc /lat/latdataquery.cgi). the results of detections made within the first 24 months of operation were released in june 2011, in the form of the second fermi-lat catalog 163 http://dx.doi.org/10.14311/app.2014.01.0163 pieter j. meintjes et al. (2fgl; see nolan et al., 2012) 1. five sources from our previous sample of unidentified egret sources have been detected with fermilat, namely 2fgl j0727.0-4726, 2fgl j1304.3-4353, 2fgl j1703.2-6217, 2fgl j1709.0-0821 and 2fgl j1815.6-6407, which are counterparts of 3eg j07244713, 3eg j1300-4406, 3eg j1659-6251, 3eg j17090828 and 3eg j1813-6419 respectively. the fermi-lat sources that coincided with the egret sources are presented as circled crosses in fig. 1. the egret (30 mev-10 gev) gamma-ray spectra of our chosen sample of sources that were observed between april 1991 october 1995 (cycles 1, 2 3 and 4 of the mission) have been determined. the photon spectral index distribution is displayed in fig. 2. the spectral distribution of these unidentified sources corresponds remarkably well with the gamma-ray blazar photon spectral index distribution observed by fermilat (abdo et al., 2010a). b=−90° b=+90° l=0° b=+10° b=−10° b=0°,l=+360° b=+30° b=+60° b=−30° b=−60° figure 1: galactic distribution of the unidentified egret sources (circles) and their fermi-lat counterparts (crosses). 1 1.5 2 2.5 3 3.5 4 3e g j0159-3603 3e g j0500+ 2529 3e g j0702-6212 3e g j0706-3837 3e g j0724-4713 3e g j0821-5814 3e g j1300-4406 3e g j1659-6251 3e g j1709-0828 3e g j1800-0146 3e g j1813-6419 3e g j1822+ 1641 3e g j1824+ 3441 γra y ph ot on i nd ex hsp-bllacs l and isp-bllacs fsrqs γ=2.03 γ=2.38 figure 2: gamma-ray photon spectral indices of the egret sources between 30 mev 10 gev. the gamma-ray spectra of the sample of egret sources with fermi-lat counterparts were determined. the spectra are presented in fig. 3. noticeable is the apparent change in the spectral index between the egret and the fermi-lat gamma-ray data, which may point to a transition in the gamma-ray production process. -100 0 100 10 100 1000 10000 100000r es id ua ls ( % ) e (mev) 1e-15 1e-14 1e-13 e 2 ( dn /d e )( w .m -2 ) 3eg j0724-4713 dn(e)=const e-2.6de a.) 3eg j0724-4713 -100 0 100 10 100 1000 10000 100000 1e+06r es id ua ls ( % ) e (mev) 1e-15 1e-14 1e-13 e 2 ( dn /d e )( w .m -2 ) 3eg j1300-4606 dn(e)=const e-3.07de c.) 3eg j1300-4406 -100 0 100 10 100 1000 10000 100000r es id ua ls ( % ) e (mev) 1e-15 1e-14 1e-13 e 2 ( dn /d e )( w .m -2 ) 3eg j1659-6251 dn(e)=const e-2.83de 1e-15 1e-14 1e-13 e 2 ( dn /d e )( w .m -2 ) 3eg j1659-6251 dn(e)=const e-2.83de b.) 3eg j1659-6251 -100 0 100 10 100 1000 10000 100000r es id ua ls ( % ) e (mev) 1e-15 1e-14 1e-13 e 2 ( dn /d e )( w .m -2 ) 3eg j1709-0828 dn(e)=const e-3de d.) 3eg j1709-0828 -100 0 100 10 100 1000 10000 100000r es id ua ls ( % ) e (mev) 1e-15 1e-14 1e-13 e 2 ( dn /d e )( w .m -2 ) 3eg j1813-6419 dn(e)=const e-2.85de e.) 3eg j1813-6419 figure 3: the gamma-ray spectra of the egret (30 mev 10 gev) and coinciding fermi-lat (20 mev 300 gev) sources. 2 gamma-ray variability the multi-wavelength emission of accretion driven sources like agn is characterized by very high luminosity, assuming isotropic emission, and variability over several time scales (e.g. fan, 2005). this is reconciled with the fact that the bulk of the spectral energy distribution (sed) of these sources is produced in a non-homogeneous and variable jet. gamma-ray flux 1http://fermi.gsfc.nasa.gov/ssc/data/access/lat/2yr catalog/ 164 modelling the multifrequency sed of agn candidates among the unidentified egret... variability has also been confirmed from agn-blazars (mattox et al., 1997). aperture photometry of the fermi-lat data over a time span of 4.6 years (from august 2008 to march 2013) has been performed to investigate possible long term variability. photons in the energy interval 100 mev to 200 gev were counted in an area with a radius of 1 degree centered on the source and monthly averages were determined. two of the sources, 2fgl j1304.3-4353 and 2fgl j1703.2-6217, showed definite signs of variability, quantified in terms of a variability index v (nolan et al., 2012), which implies that for v > 41 there is a < 1% probability of the source being steady. although variability was detected in 2fgl j1304.3-4353 (v = 47) and 2fgl j1703.2-6217 (v = 167) respectively, no periodicity could be detected (see fig. 4). 1e-07 1.5e-07 2e-07 2.5e-07 3e-07 3.5e-07 54500 55000 55500 56000 56500 p ht on f lu x (p ho to ns .c m -2 .s -1 ) time (mjd) 2fgl j1304.3-4353 (3eg j1300-4406) average a.) 2fgl j1304.3-4353 1e-07 1.5e-07 2e-07 2.5e-07 3e-07 3.5e-07 4e-07 4.5e-07 5e-07 5.5e-07 54500 55000 55500 56000 56500 p ho to n fl ux ( ph ot on s. cm -2 .s -1 ) time (mjd) 2fgl j1703.2-6217(3eg j1659-6251) average b.) 2fgl j1703.2-6217 figure 4: the gamma-ray lightcurves of the two fermi-lat sources, 2fgl j1304.3-4353 and 2fgl j1703.2-6217, which displayed variability over a period of 4.6 years. the vertical error bars indicate the 68% confidence level. 3 optical follow-up studies the optical counterparts of two unidentified egret sources have already been identified (meintjes & nkundabakura, 2012; nkundabakura & meintjes, 2012). the spectrum of pks j0820-5705 (3eg j08215814) resembles that of a fsrq at redshift z = 0.06, while the spectrum of pmn j0710-3850 (3eg j0706-3837) shows broad and narrow lines resembling the spectrum of a liner or seyfert i galaxy at redshift z = 0.129. what distinguishes the spectrum of pks j0820-5705 from that of a normal radio galaxy is the shallow k4000 depression of only 8.8 ± 2.5 %, indicating substantial non-thermal activity, while the corresponding value for pmn j0710-3850 is 80 ± 1 % (meintjes & nkundabakura, 2012; nkundabakura & meintjes, 2012), in agreement with the value expected for a liner-seyfert 1 galaxy (e.g. caccianiga et al., 1999). spectroscopic observations of other sources from our sample have been performed with the southern african large telescope (salt) (see fig. 5), equipped with the robert stobie spectrograph (rss), during 2012, in order to determine the redshift and to identify the class of agn. the spectrum of one of the unidentified egret sources, 3eg j 0159-3603, could be determined (see fig. 6), showing distinct narrow emission lines of o ii, o iii and he ii redshifted by z = 0.35. the spectrum resembles that of a typical seyfert 2 galaxy, or alternatively, a narrow line radio galaxy. this implies the possible association of two seyfert galaxies with the unidentified sources 3eg j0706-3837 and 3eg j 0159-3603 respectively. figure 5: the southern african large telescope (salt). −5e−16 0 5e−16 1e−15 1.5e−15 2e−15 4500 5000 5500 6000 6500 7000 7500 f lu x (e rg s − 1 cm − 2 å − 1 ) wavelength (å) [n ei ii ]+ h ε [o ii ] [n ei ii ] [o ii i] h γ pmn j0156−3616 (z=0.35) [o ii i] * * * * h ei i figure 6: salt spectrum of the optical counterpart of the unidentified egret source 3eg j0159-3603, showing emission lines redshifted by z=0.35 4 the spectral energy distributions (sed) the spectral energy distributions (seds) of all the counterparts corresponding to the unidentified egret 165 pieter j. meintjes et al. sources have been pieced together through multiwavelength archival data as well as multi-wavelength observations from radio to optical, using the hartebeesthoek radio astronomy observatory (hartrao) 26 m telescope, as well as various optical telescopes at the south african astronomical observatory (saao) in south africa. infrared data were obtained from the two micron all sky survey (2mass) conducted between june 1997 and february 2001. the seds of the sample of egret sources and the associated fermilat counterparts are presented in fig. 7. 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. fermi egret 3eg j0724-4713 a.) 3eg j0724-4713 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. rosat egret 3eg j1300-4406 fermi c.) 3eg j1300-4406 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. nir egret 3eg j1659-6251 fermi iras b.) 3eg j1659-6251 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. egret 3eg j1709-0828 fermi d.) 3eg j1709-0828 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. nir egret 3eg j1813-6419 fermi e.) 3eg j1813-6419 figure 7: the multifrequency seds for the sample of egret sources. the association of some of the unidentified egret sources with seyfert galaxies/liners poses a possibility of non-aligned agn and radio galaxies constituting a new class of gamma-ray sources. before discussing the modeling of the seds presented above, a brief discussion of the gamma-ray properties of some radio galaxies are presented to illustrate that these sources possess the required energetics to produce measurable γ-ray emission in the egret and fermi-lat energy domain 5 non-aligned agn: a new class of gamma-ray source at high energies (he; eγ > 100 mev) fermi-lat reported about ten misaligned radio galaxies (abdo et al., 2010b,c; rieger, 2012a). at very high energies (vhe; eγ > 100 gev), four radio galaxies have been detected, cen a (d ≈ 3.8 mpc), m87 (d ≈ 16.7 mpc) and the perseus cluster (d ≈ 77 mpc; z ≈ 0.018) sources ngc 1275 and ic 310 (e.g. rieger, 2012a). cen a was the only non-blazar detected at mev to gev energies by cgro (see steinle, 2010 for a review). the he emission from cen a reported by fermi-lat seems to come from both the extended radio lobes and the core region (abdo et al., 2010b,c). the detected he gamma-rays from the extended lobe regions suggests that particle acceleration up to vhes occurs in the disrupted jet region (bordas, 2012). the reported tev emission from cen a (aharonian et al., 2009) provided further evidence of a very effective particle accelerator in cen a. the spectrum up to 5 tev is consistent with a power-law with a photon index ∼ 2.7 ± 0.5, with no apparent variability. the he (eγ > 0.2 gev) emission in cen a is explained in terms of inverse compton (ic) upscattering of the cosmic microwave background (cmb) photons (eph = 8 × 10−4 ev) and infrared (ir) background photons by relativistic electrons with lorentz factors γe = 6 × 105, in a jet with bulk flow velocities between βγ ∼ 0.1-0.5. another peculiar aspect of cen a is a rather low, sub-eddington inferred accretion rate, ṁ ∼ 10−3ṁedd, resulting in a rather low bolometric luminosity lb ∼ 1043 erg s−1 (whysong & antonucci, 2004). the inferred equipartition magnetic field in the radio lobes is b ≈ 9 µg (rieger, 2012b), with the equipartition magnetic field near the black hole (bh) ranging between b ∼ 103 − 104 g (rieger, 2012b). the detection of non-aligned agn in the he and vhe regime poses interesting theoretical challenges regarding particle acceleration and associated gammaray emission in the jets of agn. for example, cen a shows that the radio lobes of radio galaxies may possess the required energetics to accelerate electrons to vhes producing the he gamma-rays through ic upscattering the cmb photons, even though the bulk flow lorentz factor of the jet is fairly low. the detection of tev gamma-rays is explained in terms of γe ∼ 107 electrons up-scattering disc photons to the tev domain (rieger & aharonian, 2009). the nuclear sed of cen a, based on non-simultaneous data, shows two peaks, one around ∼ 1013 hz and another around 0.1 mev (e.g. chiaberge et al., 2001; abdo et al., 2010c). the sed below a few gev seems to be satisfactorily explained by a one-zone synchrotron self-compton (ssc) model (chiaberge et al., 2001) but the same approach fail to account for the tev emission observed by h.e.s.s. (abdo et al., 2010c). 166 modelling the multifrequency sed of agn candidates among the unidentified egret... the discussion presented above underlines the fact that normal, non-aligned agn do possess the required energetics to accelerate leptons to vhe energies. the production of sub-gev gamma-rays through a ssc process in the nuclear region, combined with the he emission in the radio lobes and tev emission in the inner disc region close to the black hole (bh), presents a new paradigm in particle acceleration and gamma-ray production in agn. with this in mind, the sed of the fermi-lat counterparts of the unidentified egret sources were modelled. this is a first attempt to explain the sed of these sources within the framework of a synchrotron self-compton (ssc) and external compton (ec) model. these sources are all treated as misaligned agn, with a fixed inclination of θjet ≈ 60◦ between jet and observer. 6 sed modelling the multi-wavelength data from radio to gamma-rays have been combined to create the sed over more than 15 decades in energy (fig. 7). the egret and fermilat gamma-ray data are evaluated with the theoretical framework of a single zone synchrotron self-compton (ssc) (e.g. katarczynski, sol & kus, 2001) and external compton (ec) model (e.g. moderski et al., 2003; sikora, begelman & rees, 1994), i.e. where relativistic jet electrons up-scatter infrared (ir) photons from the disc torus and optical photons from the broad emission line (bel) regions to high energies. the model parameters for the ssc and ec models are presented in table 1 (ssc) and table 2 (ec). the corresponding sed model fits are presented in fig. 8. table 1: parameters related to ssc models for the respective sources. source 3eg j0724-4713 3eg j1659-6251 3eg j1709-0828 3eg j1813-6419 parameter units 1 radius (m) 1.5e+14 1.0e+14 1.0e+14 1.0e+14 2 b (t) 7.0e-04 7.5e-04 2.5e-04 2.5e-04 3 p1 -2 -2 -2 -2 4 p2 -3 -2.6 -2.6 -2.6 5 γmax 1.6e+03 6.2e+02 1.6e+03 2.0e+03 table 2: parameters related to the ec models for the respective sources. source r θobs γbr ke νbel αbel νir (1) (2) (3) (4) (5) (6) (7) 3eg j0724-4713 1.0e+17 1 2.0e+03 4.0e+55 10 -1 0.1 3eg j1300-4406 (bel) 1.0e+17 1 2.0e+04 6.0e+53 1 -1 0.1 3eg j1300-4406 (ir) 1.0e+17 1 3.0e+03 4.0e+54 1 -1 0.1 3eg j1659-6251 (bel) 1.0e+17 1 2.0e+03 6.0e+54 10 -1 0.1 3eg j1659-6251 (ir) 1.0e+17 1 2.0e+03 3.0e+55 10 -0.5 0.1 3eg j1709-0828 1.0e+17 1 2.0e+03 3.0e+55 10 -1 0.1 3eg j1800-0146 2.5e+16 1 5.0e+02 3.0e+55 1 -1 0.1 3eg j1813-6419 1.0e+17 1 2.4e+03 1.3e+55 10 -1 0.1 (1): radius of the emitting region (in cm), (2): viewing angle (in radians), (3): electron lorentz factor at spectral break, (4): electron normalization constant, (5): characteristic frequency of bel (in ev), (6): bel distribution photon index, (7): radiation characteristic frequency of ir (in ev). 167 pieter j. meintjes et al. 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. fermi rosat egret 3eg j0724-4713 ec (ir) scc sync scc ec (bel)ic (tot) a.) 3eg j0724-4713 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. rosat egret 3eg j1300-4406 fermi sync scc ic (tot) c.) 3eg j1300-4406 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. nir egret 3eg j1659-6251 fermi iras ec (ir) sync ssc ec (bel) b.) 3eg j1659-625 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. egret 3eg j1709-0828 fermi ec (ir) ssc sync ec (bel) ic (tot) d.) 3eg j1709-0828 1e-18 1e-17 1e-16 1e-15 1e-14 1e-13 1e-12 1e-11 1e+10 1e+15 1e+20 1e+25 νf ν (w m -2 ) ν (hz) radio opt. nir egret 3eg j1813-6419 fermi ec (ir) ssc sync ic (tot) ec (bel) e.) 3eg j1813-6419 figure 8: the modeled seds for the various egret/fermi-lat sources. from these results (see fig. 8) it can be seen that the egret and fermi-lat data are consistent with an ec model, i.e. the ic upscattering of ir photons from the disc and uv/optical photons from the line emitting clouds (broad line regions) to the egret and fermi-lat energy domain. in all the models the electron lorentz factors were γe ∼ few × 103, rather modest in comparison with the lorentz factors required to explain the γ-ray emission from, for example, cen a. 7 conclusions we report the discovery of 13 flat spectrum extragalactic sources within the error boxes of some high galactic latitude, unidentified egret sources. five of these egret sources have been detected with fermi-lat within the first 11 months of operation. in all cases the egret and fermi-lat gamma-ray emission could be successfully explained in terms of the ic upscattering of bel photons, as well as ir photons from the disc, to egret and fermi-lat energies. the adopted electron energy is of the order of γe ∼ few × 103, which is rather moderate compared to the electron energies of γe ∼ 106 − 107, inferred from the he and tev emission from cen a. these preliminary results definitely confirm that seyfert and radio galaxies could be associated with a significant fraction of the still unidentified extragalactic egret and fermi-lat sources. acknowledgement the authors thank the organisers for the invitation to present this work at this conference. some of the observations reported in this paper were performed by the southern african large telescope (salt). the salt observations were performed under the proposal: 20121-rsa-003 (pi: b van soelen). the authors thank dr. petri vaisanen from the salt science team for his guidance with the data reduction. references [1] abdo, a.a., et al.(fermi): 2010a, apj, 710, 1271 doi:10.1088/0004-637x/710/2/1271 [2] abdo, a.a., et al. 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(fermi): 2011, apj suppl., 199, 31 doi:10.1088/0067-0049/199/2/31 [16] rieger, f. & aharonian, f.a.: 2009, a&a, 506, l41 [17] rieger, f.: 2012a, in high energy gamma-ray astronomy, proc of the 5th international high energy gamma-ray astronomy meeting, heidelberg (germany), aip conference proceedings, vol. 1505, p. 80 (eds. f.a. aharonian, w. hoffman & f. rieger), publisher, melville (new york) [18] rieger, f.: 2012b, in multifrequency behaviour of high energy cosmic sources, vulcano (may 23-28, 2011), memorie s.a.it, vol 83 (1), 127 (eds. f. giovanelli & l. sabau-graziati) [19] sikora, m., begelman, m.c. & rees, m.j.: 1994, apj, 421, 153 [20] sowards-emmerd, d., romani, r.w. & michelson, p.f.: 2003, apj, 590, 109 [21] steinle, h.: 2010, pasa, 27, 431 [22] whysong, d. & antonucci, r.: 2004, apj, 602, 116 169 http://dx.doi.org/10.1111/j.1365-2966.2012.21953.x http://dx.doi.org/10.1088/0067-0049/199/2/31 introduction gamma-ray variability optical follow-up studies the spectral energy distributions (sed) non-aligned agn: a new class of gamma-ray source sed modelling conclusions 303 acta polytechnica ctu proceedings 1(1): 303–306, 2014 303 doi: 10.14311/app.2014.01.0303 atmosphere-space interactions monitor (asim): state of the art pere blay1, lola sabau-graziati2, vı́ctor reglero1, paul h. connell1, juana m. rodrigo1, juan m. macián1, josé t. biosca1, chris eyles1 1ipl, university of valencia, valencia, spain 2inta, madrid, spain corresponding author: pere.blay@uv.es abstract atmosphere-space interactions monitor (asim) mission is an esa payload which will be installed in the columbus module of the international space station (iss). asim is optimized to the observation and monitoring of luminescent phenomena in the upper atmosphere, the so called transient luminous event (tles) and terrestrial gamma ray flashes (tgfs). both tles and tgfs have been discovered recently (past two decades) and opened a new field of research in high energetic phenomena in the atmosphere. we will review the capabilities of asim and how it will help researchers to gain deeper knowledge of tgfs, tles, their inter-relationship and how they are linked to severe thunderstorms and the phenomena of lightning. keywords: space instrumentation atmosphere physics thunderstorms lightning high energy radiation. 1 introduction an increasing number of luminous phenomena related to lightning and severe thunderstorms have been discovered in the last decades. on the one hand, the transient luminous events (tles) are luminous phenomena produced at the stratosphere and mesosphere. tles studies were pushed forward with the advent of high speed cameras (capable of several tens of frames per second, see, e.g. franz et al. 1990). more recently, with very high speed cameras (thousands of frames per second, see for example cummer et al. 2006) these events have been very well characterized in terms of evolution and fine structure. on the other hand terrestrial gammaray flashes (tgfs) are very bright high energy atmospheric emissions (x-ray and γ-ray) and were discovered by astrophysical missions in orbit. according to their appearance, tles are classified mainly as blue jets, red sprites or elves. while blue jets and red sprites are directly linked to lightning and are due to air breakdown, elves are generated by electromagnetic pulses which produce ionization at around 90 km of altitude. elves have a typical lifetime of ∼1 millisecond. red sprites are generated at the mesosphere (55-80 km) and last from a few milliseconds to tens of milliseconds. blue jets are upward propagating jets of light from the cloud tops up to ∼37 km, with a velocity of ∼100 km/s and last for a few tens of a second. an schematic cartoon of these phenomena can be seen in fig. 1. with regards to tgfs, first reported detections by the batse team were published by fishman et al. (1994). tgfs are bright outburst of gamma-ray radiation produced in the atmosphere and which can be seen from space-based instruments. tgfs are thought to be produced by bremsstrahlung emission from energetic runaway electrons accelerated by electric fields on top of thunderclouds. the possible connection of tgfs to electron avalanches produced by cosmic-rays in the atmosphere, or a combination of both the cosmic-ray induced electron avalanches and the electrical discharges during lightning, is also a possibility under discussion. tgfs last typically a few hundreds of microseconds and have been observed from 40 kev up to 80 mev. for an up to date review see dwyer et al. 2012. several astrophysical high energy missions, other than batse, have detected tgfs: rhessi (designed to observe high energy emissions from the sun), fermi and agile (both dedicated to the observations of gamma-ray bursts in the universe). fermi team has even performed dedicated observations of the earth atmosphere and designed specific software and data analysis techniques in order to characterize the tgf emission. agile has been specially productive in the field of tgf detection and characterization in the 400 kev30 mev energy range. the agile team has released its own tgf catalogue available online1 (marisaldi et al. 2013). this new filed of atmospheric physics has gained 1http://www.asdc.asi.it/mcaltgfcat 303 http://dx.doi.org/10.14311/app.2014.01.0303 pere blay et al. an increasing attention from atmosphere physicist and high energy astrophysicists, and is being considered seriously by space agencies as a tool to understand atmospheric phenomena related to cloud formation, severe thunderstorms and lightning, and climate changes. in the wake of this new scientific field of research, the atmosphere-space interactions monitor (asim) was conceived as one of the first space missions dedicated and designed specifically for the observation and analysis of tgfs and tles. asim is an esa mission with contributions from spain, denmark, norway, france, italy and poland. figure 1: schematic view of tles associated to the lightning phenomena. 2 the asim mission asim was conceived as a monitoring mission with imaging and spectroscopic capabilities, with the goal to locate tgfs and tles and provide empirical data which will help to understand the nature of these emissions and the possible links between them. although tles are known to be related to lightning in severe thunderstorms, the relationship of tgfs with both of them is up to now only an hypothesis. there is some evidence that points to the relationship between tgfs and lightning, but a clear understanding is still lacking (oostgaard 2013). asim will be located at the nadir looking columbus external payload facility (cepf) on the columbus module of the international space station. asim will continuously monitor the earth atmosphere from an iss altitude of around 400 km. at a global scale, there are approximately 40 lightning flashes per second, as estimated from data taken by the optical transient detector (otd, christian et al. 1996) and lightning imaging sensor (lis, christian et al. 1999), and they are related to changes in the chemistry of the nox molecules in the atmosphere (very determinant for the ozone generation) and are thought of being of great interest for climate change studies (see for example reeve and toumi 2006, and price and rind 2012). for this reason, the observation and monitoring of lightning and related phenomenology has become of great interest for the scientific community. asim will also address this questions and provide for the first time strictly simultaneous optical and high energy data of the phenomenology related to severe thunderstorms. asim science goals will focus on understanding the physics behind tles and tgfs, get a deeper knowledge on how they are related to lightening, and understand their impact on atmospheric processes and possible links to climate determining factors. other science related to high-altitude cloud formation, cloud electrification, nox production, meteors, auroras, etc. will also be in the scope of the mission. 3 the asim instruments the asim payload will consist of two main instruments, the modular x and gamma-ray sensor (mxgs) and the modular multi-spectral imaging array (mmia). fig. 2 shows the location of the instruments on the cepa (columbus external payload adapter) which will be attached to the columbus nadir cepf. figure 2: schematic view of the mxgs and mmia instruments in the cepa. mxgs is a coded mask imaging instrument with two detector layers: one with imaging capabilities (led, low energy detector), operating in the 15−400 kev energy range, and a second layer (hed, high energy detector) sensitive to higher energy photons (200 kev−20 mev), but without imaging capabilities. the optimal 304 atmosphere-space interactions monitor (asim): state of the art energy resolution (10% at 60 kev for the led and 15% at 662 kev for the hed), the good angular location accuracy (0.7◦ for point sources and 2◦ for diffuse sources), and a time resolution bellow 5µs, make from mxgs an ideal instrument to locate tgfs and characterize their emission in terms of duration, shape of the light-curve and their spectral behavior. mxgs will have a field of view of 80◦×80◦, which at zero response becomes 147◦×147◦, i.e., the size of the earth disc as seen from the iss. table 1 summarizes the mxgs capabilities and in fig. 3 a mxgs model is depicted. table 1: summary of mxgs capabilities. led hed collecting area 1024 cm2 900 cm2 energy range 15−400 kev 200 kev−20 mev energy resolution 10% at 60 kev 15% at 662 kev time resolution ≤5 µs ≤5 µs location accuracy 0.7 (point src.) (degrees) 2 (diffuse) fov 80◦×80◦ † †147◦×147◦ at zero response mxgs may be the very first instrument locating tgfs with an error circle of ∼15 km radius, solving for the first time the ambiguities in the tgf association with lightning activity. mxgs will also be capable to observe for the very first time the tgf spectra in a very wide energy range, from 15 kev up to 20 mev. mxgs is also the very first coded mask imaging instrument specially designed to observe high energy atmospheric phenomena. figure 3: schematic view of the mxgs instrument. mmia will consist of 3 photometers and two imaging cameras. the 3 photometers are working in the wavelengths 337 nm, 180−230 nm, 777.4 nm, and the cameras are optimized to work at 337 nm and 777.4 nm, covering altogether from the near uv up to the near ir. the time resolution is 83 ms for the cameras and 0.01−0.1 s for the photometers. the cameras will have an spatial resolution between 0.4 and 0.5 km and both the photometers and the cameras have the same field of view, namely, 61.4◦. mmia will provide invaluable data on the optical emission (as opposed to the high energy emission) from severe thunderstorms. table 2 summarizes the capabilities of mmia and in fig. 4 a model of the instrument is depicted. figure 4: schematic view of the mmia instrument. table 2: summary of mmia capabilities. cameras photometers fov (nadir) 61.4◦ 61.4◦ num. of pixels 1024×1024 −− pixel resolution 0.4−0.5 −− (nadir) (km) time resolution (ms) 83 0.01−0.1 spectral bands 337/5 337/5 (center/width) 777.4/5 777.5/≤5 (nm) 205/50 mxgs and mmia will both work together in their triggering mode, and will be able to send their triggering information to each other. fig. 5 depicts a schematic view of the instruments capabilities and their interconnected cross-triggering system. they will offer the first strictly simultaneous observation of luminous emissions associated to thunderstorm activity. 305 pere blay et al. figure 5: schematic view of the mmia and mxgs configurations, depicting the instruments capabilities and the interconnected cross-triggering. 4 mission status asim mission was approved by esa during 2005. now (2013) the critical design review (cdr) is on going and the manufacturing is expected to start along 2013. the flight model should be ready by the end of 2014 and the launch of the mission is expected by mid-2015. the nominal operational life of the mission is 2 years, with possible extensions. 5 conclusions asim will help the scientific community to solve some of the problems related to tle and tgf formation, a quite recent new filed of atmospheric physics which has been also called high energy atmosphere physics (see dwyer et al. 2012). the capabilities of its instruments and the wavelength coverage will allow the location tgfs with high accuracy and disentangle their links to tles, thunderstorms and lightning. acknowledgement pere blay acknowledges funding from the spanish ministerio de economia y competitividad through project aya-2011-29936-c05. references [1] christian h.j., driscoll k., goodman s.j., blakeslee r.j., mach d., and buechler d. 1996, proc. 10th int. conf. on atmospheric electricity, osaka, japan, icae, 368371 [2] christian h.j., blakeslee r.j., goodman s.j., et al. 1999, proc. 11th int. conf. on atmospheric electricity, guntersville, al, international commission on atmospheric electricity, 746749 [3] cummer, s. a., n. jaugey, j. li, w. a. lyons, t. e. nelson, and e. a. gerken 2006, geophys. res. lett., 33, l04104 doi:10.1029/2005gl024969 [4] dwyer j.r., smith d.m., and cummer s.a. 2012, space sci rev, 173, 133 doi:10.1007/s11214-012-9894-0 [5] fishman g.j., bhat p.n., mallozzi r. et al. 1994, science, 264, 1313 doi:10.1126/science.264.5163.1313 [6] franz r.c., nemzak r.j., and winkler j.r. 1990, science, 249, 4851, 1990 [7] m. marisaldi, f. fuschino1, m. tavani et al. 2013, j. geophys. res. space physics, http://dx.doi.org/10.1002/2013ja019301 [8] n. ostgaard, t. gjesteland, b. e. carlson, a. b. collier, s. a. cummer, g. lu and h. j. christian 2013, geophys. res. lett.,vol. 40, 1-4 doi:10.1029/2012gl054022 306 http://dx.doi.org/10.1029/2005gl024969 http://dx.doi.org/10.1007/s11214-012-9894-0 http://dx.doi.org/10.1126/science.264.5163.1313 http://dx.doi.org/10.1029/2012gl054022 introduction the asim mission the asim instruments mission status conclusions acta polytechnica ctu proceedings doi:10.14311/app.2015.1.0022 acta polytechnica ctu proceedings 2:22–28, 2015 © czech technical university in prague, 2015 available online at http://ojs.cvut.cz/ojs/index.php/app vector maps in mobile robotics ales jelinek central european institute of technology, brno university of technology, technicka 10, brno, czech republic correspondence: alesjelinek@ceitec.vutbr.cz abstract. the aim of this paper is to provide a brief overview of vector map techniques used in mobile robotics and to present current state of the research in this field at the brno university of technology. vector maps are described as a part of the simultaneous localization and mapping (slam) problem in the environment without artificial landmarks or global navigation system. the paper describes algorithms from data acquisition to map building but particular emphasis is put on segmentation, line extraction and scan matching algorithms. all significant algorithms are illustrated with experimental results. keywords: robotics, slam, vector map, point cloud, edge extraction. 1. introduction mapping in robotics is a subject of research for a number of years and large progress was already achieved. paper [1] describes theoretical way, how to build and update a consistent map, which converges to a precise image of the real environment (at certain level of detail). through the years, more improvements and new algorithms were invented for this task, good tutorial papers are for example [2] [3], or the book [4]. formulation and mathematical description of the slam problem is unified and generally accepted today, but a complete solution for 3d dynamic environment is still not known and slam as a whole is still under active research. the most of today’s work in the field is concentrated on two main bottlenecks of known algorithms: (1.) computational optimization. every map contains some data describing the real environment. computational efficiency is usually evaluated in big o notation with respect to n, reflecting the number of cells in an evidence-grid map, the number of landmarks in a geometric map, etc. non-optimized solutions have o(n2) or even exponential complexity, but several algorithms are more efficient, for example fastslam [5] with o(m log(n)) complexity (where n is the number of landmarks and m the number of particles in the rao-blackwellized filter). this algorithm can handle orders of magnitude more landmarks than o(n2) solutions. computational efficiency is a key to large maps with a lot of details. (2.) feature recognition and matching. robust and reliable feature (landmark) extracting and matching is crucial for most of slam algorithms. new data has to be associated with older measurements precisely, or at least with negligible number of mismatches. feature extraction is needed when camera or laser scanner is used to provide information about the robot surroundings. for example corners and t-junctions are usually searched in input data and then the whole pattern is correlated with an already known map. several complications may arise, because the same feature may look different from different points of view and data registration in its straightforward implementation is exponentially complex task. data structure of a map is closely related to both of these problems. operations executed on the map have to be as fast as possible especially updates with the new information and feature search/comparison are crucial. representation of obstacles, free space and unexplored areas effects feature recognition robustness. another attribute of a map is its ability to approximate the real environment. importance of this attribute depends on further usage of the map. if the robot is meant to explore unknown area and provide information about it, requirements will be probably higher than in case of a robot, which just needs to avoid obstacles and memorize a short history of its path. there are more possible ways to represent a robotic map. one of the most common is robot evidence grid, which parcels continuous world into a set of squares (cubes in 3d). size of a square is optional and may be even adaptive to better fit the reality, but this style of approximation of smooth world is still rather crude, because straight edges which are not co-linear with grid lines are jagged regardless of grid resolution. evidence grids are widely used and well established in many applications, where this downside is negligible. visually better results are achieved with vector maps, where edges are represented with line segments. good example of classical vector based mapping system is vecslam described in [6]. the aim of authors work is to create a system for high quality mapping, therefore vector maps were chosen as a good approximation of the real world. map representation is only one part of the problem. robot exploring unknown environment needs 22 http://dx.doi.org/10.14311/app.2015.1.0022 http://ojs.cvut.cz/ojs/index.php/app vol. 2/2015 vector maps in mobile robotics to update its map with new information. data are usually obtained from rangefinder sensors such as laser scanners and sonars in the form of a point cloud and need to be incorporated in the current map. probably the most used technique for point cloud registration is iterative closest point (icp). good description and historical survey of this method is available in [7]. efficient variants of this method were published in [8]. furthermore, it is possible to use icp for registration of more complex formations, such as surfaces of scanned objects. this approach is described in [9] and [10]. there are several alternatives to icp, which are trying to generalize the notion of “closeness” or “similarity” from points to lines, triangles etc. similarity of points is usually measured as their euclidean distance. an edge is often described as a segment of a line. having two such segments, we can define an area they demarcate and use it as a measure of similarity. this definition is not unified and it is possible to find several different approaches, for example sum of squared distances from point in a scan to the nearest edge in the map [11], penalizing function of “non collinearity” and spatial distance [12], or a criterion based on newly introduced matrix scalar product [13]. another interesting approach is critical ray method proposed in [14], which does not use line similarity at all. diversity of these approaches is high and none of them is considered to be remarkably better than the others, therefore the author have decided to develop his own method and test new ideas, because this field of research is definitely not fully explored. reducing point cloud to a set of line segments (in 2d) or triangles (in 3d) has very useful properties. at first, information from hundreds or thousands of points is expressed with (at most) tens of objects. data reduced in count are processed faster and therefore more complex algorithms may be used. on the other hand, mathematics becomes more complicated, because lines and faces have more degrees of freedom and constraints than points and as written above, similarity is not as easy to enumerate. the process of line extraction has generalizing ability. after interpolation by line segments (or triangles), corners become sharper and noise is suppressed, which results in better looking (more corresponding to the real world) map. in addition (depending on line extraction algorithm) edges might be estimated with higher accuracy than single point due to averaging effect. there is no restriction on using n − 1 dimensional objects in n dimensional space. a generalization of icp for lines in 3d space is described in [15]. paper [16] describes whole set of algorithms for matching sets of lines in 3d and many others are possible to find. for even better approximation of the real world, curves and curved planes could be used. these are more complicated for computation and not too much papers are related to this topic. an example of such approach is a method [17] based on graph theory. 2. point cloud processing system the overall diagram of the measurement and mapping process is depicted in fig. 1. every time a measurement is acquired, the whole process is run through. each rectangular box in the diagram means one algorithm performing operation on its input data. the whole process leads to update of the map the robot creates and the pose of the robot itself. slam itself is implemented in a very elementary way and no attempts to build converging map were made yet. the main objective in current state of development is to test data processing, edge extraction and scan registration algorithms. laser scanner used to obtain all data in this chapter is velodyne 32 hdl, which provides approximately two thousand points (per turn) in polar coordinates for each of its thirty two lasers. only horizontal set of measurements was taken into account for 2d map. at the beginning, all data from sensors need to be converted into the data format useful for following algorithms. this step also ensures usability of the system for any instrumentation with suitable properties. a disadvantage of this process is a certain enlargement of data size, because laser scanners and other sensors usually use some sort of data compression to minimize data flow. in case of velodyne 32 hdl this means conversion from 16-bit integer to 32-bit floating point number. to speed up corresponding edge selection, crude pose estimation is useful. it enables to specify part of the map, where the robot most probably is, and therefore, the search space of correspondences is reduced a lot. as is well known in robotics community, techniques such as inertial sensors or odometry cause integration of error of pose estimation. this makes them useless after a certain amount of time. in the presented algorithm, these are only used for measurement of relative pose change between two following scans. this ensures that the estimated pose error is always in a known limit. to get an absolute pose, this estimation is added to the last known pose, determined in the previous iteration of the data processing (dashed arrow leading to “pose estimation” block in fig. 1). 2.1. filtering and clustering treatment of raw point cloud is a more complex task. at first it is necessary to filter out noise and outliers and to find dense clusters of points defining edges of solid objects in the environment. raw data from one set of measurements are shown in fig. 2a. it is clear that scans are full of irrelevant points and odometry was not very accurate. some noise is always present. laser scanner measures distance with certain precision, which is stated in the datasheet. the other sake of noise is too dissected environment, where individual faces of objects 23 ales jelinek acta polytechnica ctu proceedings measured data conversion from scanner data format to cartesian and polar coordinate system point cloud segmentation and filtering edge extraction corresponding edge selection finding the best transformation for incorporation of the scan in a map map update map inertial sensors pose estimation most probable pose pose correction data flow data from previous step scanner data processing operation result conversion form scanner data format to common physical quantities figure 1. overall scheme of point cloud processing system from data acquisition to map and pose update. are so small that rangefinder cannot scan them properly. a good examples are treetops and bushes. noisy parts of point cloud are best to treat as a solid body with defined borders and density. other possibility is to filter them out, which causes loss of information, but on the other hand, with a careful filtering, it is possible to “see” through low density objects (e.g., see wall behind a bush). outliers in point cloud are usually caused by scanning too distant or too close objects, which are out of the range of the particular laser scanner. these are easy to identify and can be removed without serious impact on the quality of a scan. more interesting to deal with are outliers originating from a reflection. in a common environment, there can always be objects, which are able to reflect laser beam and corrupt the measurement. if the reflecting object is small and curved, few beams, which hit it, are scattered wide apart and appear as randomly distributed single points in a point cloud. these are generally easy to remove. another case are large, flat objects such as mirrors and windows, which uniformly reflect a large amount of rangefinder beams. this effect causes phantom objects to appear in the scan and there is probably no way to detect this without additional sensors or a set of scans from different positions. clustering is used to remove all unnecessary points from a scan and to determine dense clusters of points, which could potentially form an edge of a real object. an algorithm used to get result depicted in fig. 2b is as follows: from the laser scanner we get data sorted by an angle. algorithm proceeds from one point to another and checks if k previous points are closer or farther, than a certain distance. three alternatives may occur. a point might be too far away from all of the previous ones. this leads to establishment of a new cluster. if the working point is close enough to some previous points and all these points belong to the same cluster, the working point is added to it. third case arises, when the working point can belong to more than one cluster. in such a situation, all those clusters and the working point are joined together. in the end, clusters with the size under a certain threshold are called outliers and are removed from the scan. from a practical point of view, it is wise to make the distance comparison threshold adaptive, because the farther the detected obstacle is, the sparser the measured points are. scaling the threshold linearly with the distance between the working point and the robots position is a good strategy. the proposed algorithm safely removes noise from too dissected objects as well as random outliers, because these create large number of very small clusters (typically less than ten points), which are removed in the final stage of the process. phantom objects are not recognized. noise caused by laser scanner itself removes the following algorithm for line extraction. this theme is directly related to line simplification algorithms (because points in a scan are sorted) and both of these are discussed in a lot of fields of re24 vol. 2/2015 vector maps in mobile robotics (a) . raw data. (b) . segmentation. (c) . line extraction. (d) . complete map. figure 2. a) depiction of all measured point clouds (yellow) with estimated robot path (green). no scan fitting or other postprocessing is used, image corresponds to raw data from sensors. b) segmentation of one scan is shown on this figure. each color means a cluster, which is believed to be continuous edge in reality and which is separate from the others. noise and outliers are filtered out and not displayed. c) line extraction in segmented point cloud (black). each cluster is approximated by a line or a poly-line (red). d) a map composed using the line similarity criterion. red edges are long enough to participate in scan fitting process, blue edges are shorter and just add details to the map. green trajectory is acquired from odometry and dark blue one is an estimation of true trajectory. search. in addition to robotics, similar approaches can be found in cartography, computer vision and computer graphics. plenty of algorithms were developed during last decades, having different properties, such as speed, memory requirements and quality of approximation. good survey comparing line extraction/simplification algorithms in robotics is [18]. results of this paper clearly show that algorithms such as ransac or hough transform are not fast enough for online processing of point clouds. good performance was observed when using splitmerge, incremental and line regression algorithms. split-merge (also known as douglas-peucker algorithm [19]) was the fastest one and belongs to wide family of o(n) and o(n log(n)) complex algorithms used in cartography [20]. main drawback of these algorithms is that they are based on point elimination, and therefore, a lot of information is discarded. the edge extraction algorithm from the diagram in fig. 1 is based on line regression and works as follows: it passes through the cluster and computes the least square approximation and the variance of selected set of points. if the variance gets over predefined limit, the line is saved, and algorithm proceeds computing new line. the last line terminates at the end of the cluster. this approach is advantageous, because all points contributes to the result and contained information is better utilized. although loss of information was overcome by the least squares approximation, the algorithm is still not optimal in terms of “best fit poly-line”. all but the last line have maximal permitted variance, which means, that for too high threshold lines go over corners and an approximation is worse, than it could be. more technically, the total sum of squared distances from each point to its regression line is higher, than is achievable with given number of lines. fine tuning of the threshold is therefore necessary. complexity of the algorithm is o(n). result of line extraction is depicted in fig. 2c. 2.2. scan fitting when updating a map with a new scan, corresponding edges must be selected at first. it is responsibility of an operator or a path planing algorithm to obtain new scans frequently enough to ensure overlap of a new scan and the map. this means the robot always has to see part of the already known environment (except the first scan of course). acquiring more scans from one location leads to a more precise map (averaging effect). scan fitting is the point, where crude pose estimation comes in useful, because searching through the whole map would be very computationally expensive. knowing approximate pose and its maximal error allows us to considerably reduce search space and keep computational complexity constant, independent on the size of the whole map (assuming the number of visible details is comparable across the map). if the robot is suddenly moved far away from its previous location, it becomes lost and the algorithm is not capable of working correctly anymore. only long enough edges are used for the process, because the shorter ones are not determined with a sufficient precision. the maximal angular error of the pose estimation is directly used as the maximal angle between possibly similar edges. if the angle is inside a tolerance, shortest distances of end points of first edge to the second edge are computed. if at least 25 ales jelinek acta polytechnica ctu proceedings one is shorter than maximal permitted shift between two scans, edges are overlapping each other and can belong to one real edge. there are three possible situations after this step. for some edges from the new scan, there is no similar one in the map. these probably belong to an area which was not explored yet, or where only short edges were determined. new edges do not contribute to scan fitting process. the second group of new edges has one similar match in the map and these are directly used for scan matching. in the last group, there are edges with more than one possible match. at this situation, it is hard to say, which possible pair is the best one, or if more edges should be connected together. current practice is to remove edges already used in the second group and from the rest to select the closest one (in the sense of overlapping). in rare cases, the edges corresponding in real world are not the closest ones. this leads to mismatch and can even corrupt the map, but in most cases, correctly matched edges prevail and the mismatch is unnoticeable. at this point, pairs of corresponding edges from the new scan and the map are known and the actual fitting may take place. as well as in the previous step, the line similarity criterion will be used, but now with slightly different properties. an idea, that lines are more similar as the angle they form is smaller is still true, but overlapping criterion used before is not advantageous. instead, as already mentioned in the introduction, an area defined by both edges is used. the main idea of the criterion is depicted in fig. 3. let two edges be defined by lines p and q with the end points ab and cd. at first, an axis of angle formed by these edges is found. then, for all end points, their perpendicular projections (with respect to the axis o) on an opposite line are found. two end points and their projections form a quadrangle as shown in fig. 3 for points ab (both cases). there are six possible quadrangles and the largest one has to be used for the criterion enumeration. the square of that area is used as a similarity metric. the function for the similarity enumeration is not strictly convex, therefore for one pair of edges, there is not a single minimum defining one pose in which both edges are the most similar. instead, there is an infinite number of possible poses, in which the criterion is zero. this property is clear from the geometrical visualization of collinear edges sliding over each other and corresponds to the problem of localization in a long corridor. if a robot “sees” only collinear walls, it is not able to determine the displacement along an axis of the corridor one degree of freedom cannot be estimated from available data. for an exact solution, at least two skew pairs of edges must be used. fitting itself is an iterative process, because no closed form solution was found so far. a new scan is transformed using the standard homogeneous transformation. the goal is to find such translation and (a) . no intersection. (b) . with intersection. figure 3. line segments similarity criterion. the smaller is the gray area, the more similar are both line segments. rotation, so that the sum of the squared areas of all edge pairs is the smallest possible. nelder-mead simplex method was used to find the global minimum. for longer edges, the area they demarcate is large even if an angle they form is relatively small. this means, that longer edges have greater impact on scan fitting, which is correct, because long edges are determined by more points, and therefore, be more precisely known. 2.3. map building after the fitting process is finished, a new scan is transformed using the found parameters and edge pairs are joined to form a new set of edges in updated map. the same transformation, which was used to update the map, is applied on the path estimate to correct the odometry error. assuming static environment, the information about pose gotten from scan fitting is absolute, and therefore, an error should not have integrative characteristics, however, this is ideal state which requires a precise map corresponding to the environment. the main problem of map building process in current state of development is that the map does not converge. scans are incrementally fitted to the map, but new measurement influences only the part of the map, which is currently visible and new information is not propagated through the whole map to assure convergence. in other words, change in length of one edge should affect all connected edges (including those which are not actually seen), but current implementation is not able to handle this. integrative 26 vol. 2/2015 vector maps in mobile robotics figure 4. a larger map composed using the presented processing system. edges are colored in the same way as in fig. 2d. gray trajectory is acquired from odometry and dark blue one is an estimation of the true trajectory. characteristics of error of pose and map is therefore not suppressed. on the other hand, using edges and line similarity criterion for scan fitting is very accurate and divergence process is much slower. results of this step are maps in fig. 2d and fig. 4. in comparison with a naive scan merging in fig. 2a, new maps are much more usable for navigation as well as environment documentation. further processing for even better results can follow at this point, one example of possible operations is additional merging of shorter edges (blue edges in fig. 2d and fig. 4) to form a cleaner map. 2.4. experimental results practical testing was made in office and laboratory environment. algorithms based on edge detection are well suited for such application, because artificial objects are usually simply shaped with a lot of flat surfaces. on the contrary, a natural environment is much more complicated with lot of details and lack of flat surfaces, therefore effective line extraction would be very complicated or even impossible. the robot used for testing was orpheus x3, which is four-wheeled vehicle with differential steering. laser scanner used was velodyne 32 hdl. during the experiments, the robot was manually operated and all measurements were taken, when the robot was not moving. the first experiment used for demonstration of the processing pipeline is depicted in fig. 2a to fig. 2d. ten measurements 60 centimeters apart were taken. the odometry error was intentionally increased to demonstrate correction ability of scan fitting algorithm. outliers and noise filtration was also proved to be working, for example small cloud of points on top of fig. 2a was correctly removed. the second experiment was made in a larger scale on a corridor and resulting map is shown in fig. 4. robot moved from starting position to the end of the corridor and then returned back. total length of the traveled path was 46.710 meters and 33 measurements were taken along the way. the difference between the real and the estimated position of the last point of trajectory was only 6 mm, which is even less than the precision of used laser scanner (20 mm). the experiment was held only once, therefore repeatability is not yet determined. fairly good results are caused by two main factors. at first, the line extraction algorithm with the least squares approximation is capable of reducing noise through averaging and provide more accurate edge positions than single points in a point cloud. the second reason is that the robot moves in one room and always sees edges from the first scan, therefore an error of map and pose is not integrating. if the robot would move along a path with a lot of turns and after some time would not see its starting location, this problem would start to manifest. however, these experiments were made to proof the concept of scan fitting via edge similarity criterion and slam itself is not yet at the center of the research. 3. conclusion and future work the paper presents the set of algorithms used to process data from laser scanner and build a vector based map. the whole scheme of the process is depicted in fig. 1. most algorithms used are new and still under testing, but findings stated in literature (e.g., [10], [12], [16]) were proved to be true, because scan matching based on edge detection yields quality results and definitely can be used for solving slam problem. at least, considering results in fig. 2d and fig. 4, current state of the research is very promising. the main task for the nearest future is development of good representation of the topological-metric map, to enable a real slam to be implemented. topological information is necessary for visible (or just close) edges finding algorithm and for path planning. metric information is crucial for mapping itself, scan matching, and path planning. term “good representation” means such data structure, that would be efficient for search and computing algorithms. nearest neighbor of polygon search or finding an edge in area with certain coordinates are only a few examples of problems which are necessary to be efficient in a good mapping algorithm. the second requirement expected from a good map is a convergence. as the robot explores surrounding 27 ales jelinek acta polytechnica ctu proceedings environment, the map should adapt to be closer to reality. this is not achieved now and lot of work is focused on this task. with an authentic map and well working fitting algorithm, localization itself should not be too complicated and a solution for the slam problem using vector maps should be completed (at least in its main aspects). acknowledgements this work was supported by the project ceitec central european institute of technology (cz.1.05/1.1.00/02.0068) financed from european regional development fund. references [1] m. dissanayake, p. newman, s. clark, et al. a solution to the simultaneous localization and map building (slam) problem. ieee transactions on robotics and automation 17(3):229–241, 2001. doi:10.1109/70.938381. [2] h. durrant-whyte, t. bailey. simultaneous localization and mapping: part i. ieee robotics & automation magazine 13(2):99–110, 2006. doi:10.1109/mra.2006.1638022. [3] t. bailey, h. durrant-whyte. simultaneous localization and mapping (slam): part ii. ieee robotics & automation magazine 13(3):108–117, 2006. doi:10.1109/mra.2006.1678144. [4] s. thrun, w. burgard, d. fox. probabilistic robotics. the mit press, 2005. [5] m. montemerlo, s. thrun, d. koller, b. wegbreit. fastslam: a factored solution to the simultaneous localization and mapping problem. aaai/iaai 2002. [6] h. j. sohn, b. k. kim. vecslam: an efficient vectorbased slam algorithm for indoor environments. journal of intelligent and robotic systems 56(3):301–318, 2009. doi:10.1007/s10846-009-9313-2. [7] l. oswald. recent development of the iterative closest point algorithm, 2010. [8] s. rusinkiewicz, m. levoy. efficient variants of the icp algorithm. proceedings third international conference on 3-d digital imaging and modeling pp. 145–152. doi:10.1109/im.2001.924423. [9] p. besl, h. mckay. a method for registration of 3-d shapes. ieee transactions on pattern analysis and machine intelligence 14(2):239–256, 1992. doi:10.1109/34.121791. [10] z. zhang. iterative point matching for registration of free-form curves and surfaces. international journal of computer vision 13(2):119–152, 1994. [11] h. pottmann, s. leopoldseder, m. hofer. registration without icp. computer vision and image understanding 95(1):54–71, 2004. doi:10.1016/j.cviu.2004.04.002. [12] j. elseberg, r. t. creed, r. lakaemper. a line segment based system for 2d global mapping. 2010 ieee international conference on robotics and automation pp. 3924–3931, 2010. doi:10.1109/robot.2010.5509138. [13] q. li, j. griffiths. iterative closest geometric objects registration. computers & mathematics with applications 40(10-11):1171–1188, 2000. doi:10.1016/s0898-1221(00)00230-3. [14] e. tsardoulias, l. petrou. critical rays scan match slam. journal of intelligent & robotic systems 72(3-4):441–462, 2013. doi:10.1007/s10846-012-9811-5. [15] m. alshawa. icl : iterative closest line a novel point cloud registration algorithm based on linear features. ekscentar (10):53–59, 2007. [16] b. kamgar-parsi, b. kamgar-parsi. algorithms for matching 3d line sets. ieee transactions on pattern analysis and machine intelligence 26(5):582–93, 2004. doi:10.1109/tpami.2004.1273930. [17] k. demarsin, d. vanderstraeten, t. volodine, d. roose. detection of closed sharp edges in point clouds using normal estimation and graph theory. computer-aided design 39(4):276–283, 2007. doi:10.1016/j.cad.2006.12.005. [18] v. nguyen, s. gächter, a. martinelli, et al. a comparison of line extraction algorithms using 2d range data for indoor mobile robotics. autonomous robots 23(2):97–111, 2007. doi:10.1007/s10514-007-9034-y. [19] d. h. douglas, t. k. peucker. algorithms for the reduction of the number of points required to represent a diditized line or its caricature. cartographica: the international journal for geographic information and geovisualization 10(2):112–122, 1973. doi:10.3138/fm57-6770-u75u-7727. [20] w. shi, c. cheung. performance evaluation of line simplification algorithms for vector generalization. the cartographic journal 43(1):27–44, 2006. doi:10.1179/000870406x93490. 28 http://dx.doi.org/10.1109/70.938381 http://dx.doi.org/10.1109/mra.2006.1638022 http://dx.doi.org/10.1109/mra.2006.1678144 http://dx.doi.org/10.1007/s10846-009-9313-2 http://dx.doi.org/10.1109/im.2001.924423 http://dx.doi.org/10.1109/34.121791 http://dx.doi.org/10.1016/j.cviu.2004.04.002 http://dx.doi.org/10.1109/robot.2010.5509138 http://dx.doi.org/10.1016/s0898-1221(00)00230-3 http://dx.doi.org/10.1007/s10846-012-9811-5 http://dx.doi.org/10.1109/tpami.2004.1273930 http://dx.doi.org/10.1016/j.cad.2006.12.005 http://dx.doi.org/10.1007/s10514-007-9034-y http://dx.doi.org/10.3138/fm57-6770-u75u-7727 http://dx.doi.org/10.1179/000870406x93490 acta polytechnica ctu proceedings 2:22–28, 2015 1 introduction 2 point cloud processing system 2.1 filtering and clustering 2.2 scan fitting 2.3 map building 2.4 experimental results 3 conclusion and future work acknowledgements references 255 acta polytechnica ctu proceedings 1(1): 255–258, 2014 255 doi: 10.14311/app.2014.01.0255 probing gravitational theories with eccentric eclipsing detached binary stars leopoldo milano1,2, rosario de rosa1,2, mariafelicia de laurentis1,2, fabio garufi1,2 1dipartimento di fisica universita’ federico ii di napoli complesso universitario di monte s. angelo, via cinthia, i 80125, napoli, italy 2sezione infn di napoli,complesso universitario di monte s. angelo, via i 80125, napoli, italy corresponding author: milano@na.infn.it abstract in this paper, we compare the effects of different theories of gravitation on the apsidal motion of eccentric eclipsing detached binary stars. the comparison is performed by using the formalism of the post-newtonian parametrization to calculate the theoretical advance at periastron and compare it to the observed one, after having considered the effects of the structure and rotation of the involved stars. a variance analysis on the results of this comparison shows that no significant difference can be found due to the effect of the different theories under test with respect to the standard general relativity (gr). it will be possible to observe differences, as we would expect, by checking the observed period variation on a much larger lapse of time. keywords: relativity gravitation eclipsing binaries apsidal motion. 1 introduction the problem of the motion of two bodies under their mutual gravitational attraction and the study of binary stellar systems has always been the ideal test bed for the theories of gravitation. several authors in the last decades dedicated a lot of work in analyzing, both on the theoretical and the experimental point of wiew, the phenomenon of the periastron precession in binary systems to test various gravitational theories [1, 2] as well as to find correction to the newtonian and general relativistic behaviour of the systems due to stellar form factors, spin, tides and other phenomena [3]. the classical effect of general relativity (gr) on the apsidal motion rate at periastron is well known since long time and described by levi-civita in a famous paper in 1937 [4, 5]. another possible formulation of the problem, that allows also to test other gravitational theories besides gr, is the use of parametrized post newtonian (ppn) formalism [6, 7]. using this formalism, the different gravitational theories can be compared side by side on the basis of a set of post-newtonian (pn) parameters: the masses, the system major semiaxis and the eccentricity. thus, using a sample of eccentric eclipsing detached binary (eedb) systems, for which masses and orbital parameters are known with sufficient precision, it is possible to compare the apsidal motion rate at periastron ω̇t h, as expected in the different theories with the observations in order to verify whether the observations can select one theory or another. the choice of this class of stellar objects to test the theories is dictated by the circumstance that the both orbital and the structural parameters can be precisely determined by observation of the eclypses (the only alea being the orbit plane observation angle), and that in close stellar orbits, the gravitational field is supposed to be strong, thus enhancing eventual effects of deviations from the classical theory. there are so many aspects of binary star evolution and angular momentum exchange (see, e.g., [8]) that any attempt to dig out from these other effects the subtleties of gr could seem futile, nonetheless, we consider useful to ascertain whether from the observation of many binary systems, some statistical information about the prevalence of one or another theory of gravitation can be extracted and to ground the bases of a method that can be used when other more significant data will be available. 2 advance at periastron the idea of considering relativistic gravitational tests in terms of a metric expansion is originally based on a work by shiff [9] who expanded the single body metric in terms of the ratio between the geometrized mass mg = gm/c 2 and the distance r: g00 = 1 − 2α mg r + 2β (mg r )2 255 http://dx.doi.org/10.14311/app.2014.01.0255 leopoldo milano et al. g0k = 0 gik = − ( 1 + 2γ mg r ) δik i,k = 1, 2, 3 (1) four new parameters α′, α′′, α′′′, ∆ were then introduced to account for relative velocities and accelerations. for general relativity all the parameters are equal to 1 and the advance at periastron can be verified to reproduce the ”classical” formula by levi civita [4]. we conveniently modified the ’classical’ formula by introducing a factor (kt h) to take into account the dependance on the theory dependent pn terms in order to test the different relativistic theories. [10]: ω̇rel = 1.8167 × 10−4kt h ( m p ) 2 3 c2(1 −e2) (2) where: kt h = αt h(8∆t h −αt h + 2α ′′ t h −α ′′′ t h −γt h −α ′ t h) 2 (3) the pn parameters, are calculated for the different gravitational theories , i.e. the general relativity (’classical’ term), the nordvedt, the brans-dicke theories and the so called f(r) theories that take into account higher order terms of the ricci scalar r, giving a general expression for the relativistic term ω̇rel, that contributes to the advance at periastron. using for the different theories the appropriate values of αt h,α ′ t h,α ′′ t h,α ′′′ t h, ∆t h and γt h the numerical values of kt h can be obtained [10]: kt h =   kgr = 3 → (general−relativity) kbd = 19 7 → (brans−dicke) knd = 11 4 → (nordvedt) kf(r) = 13 4 → (f(r)) (4) 3 comparison with experimental data to test the effects of deviations from the gr, we choose to study binary stellar systems with small radius orbits, so that the gravitational field is strong enough to evidence these deviations, if any. among the various binary stars catalogues available in literature, we choose a sample of eccentric eclipsing detached binary stars such that the period, the eccentricity, the masses of the components, and, possibly, the observed internal structure function are known with a good precision as e.g. [11]. for these systems the passage at periastron precedes in a way that is precisely predictable from the gravitational theory, once given the stellar parameters such as masses, radius of the components, and orbital parameters. to compare the global rate of theoretical apsidal motion in a binary system with the measured one we must take into account the individual contributions of each component due to tidal and rotational distortions, and the general relativistic term ω̇th, where the index th indicates the theory under test (e.g. ωgr for general relativity). assuming that rotation of both components of an eclipsing binary system is perpendicular to the orbital plane, the apsidal motion rate, ω̇ is given by the following simple relation [12]: ω̇obs = ω̇cl + ω̇rel (5) where ω̇cl is the classical newtonian term and ω̇rel is the relativistic contribution of eq. 2 and eq. 4. the dependance of ω̇cl on the binary system parameters is expressed through the internal second-order constants (isc): k̄2t h = c21k21t h + c22k22t h c21 + c22 (6) k̄2obs = ω̇cl c21 + c22 (7) where the parameters c2i are related to the masses and the orbital eccentricity of the binary system. it must be noticed that the individual isc’s k2,i cannot be obtained from the observations although they can be interpolated from evolutionary codes like those used in [13, 14]. so we can evaluate a mean model dependent k̄2t h and a mean observation dependent k̄2obs, and compare them to test the evolution stellar models from the observations of apsidal motion. taking into account that k̄2obs is generally smaller than k̄2t h (this means that the evolution models predict stellar cores less dense than those found by observed data), remembering eq. 2 and eq. 5 we see that ω̇cl will vary according to ω̇t h. in this way we can test the different relativistic theories of gravitation by verifying whether the agreement of the model dependent mean isc, with those coming from observations is significantly improved by using the different ω̇t h relativistic terms. in fig. 1 we show the trend of the apsidal motion rate ω̇gr vs ω̇gr,bd,nd,f(r). it results: ω̇bd ∼= 0.92ω̇gr, ω̇nd ∼= 0.90ω̇gr, ω̇f(r) ∼= 1.10ω̇gr. obviously the numerical coefficients are the ratios kt h kgr (see eq. 4). it is interesting to note that the f(r) theory gives a relativistic contribution that is slightly higher than gr,bd and nd. moreover, gr is ≈ the mean between f(r) and bd and nd theories. it is also evident that there is no significant difference among the theories under test within the errors. in fig. 2 we show the trend of observed ω̇obs vs ω̇rel for different relativistic terms (th ≡ gr,bd,nd,f(r)). the red line is the trend of ω̇rel = ω̇obs. it is evident that, apart 256 probing gravitational theories with eccentric eclipsing detached binary stars from a few systems, the relativistic term is always less than the observed one. 0 0.5 1 1.5 2 2.5 3 3.5 x 10 −3 0 0.5 1 1.5 2 2.5 3 3.5 x 10 −3 dω/dtgr d ω /d t g r ,b d ,n d ,f (r ) dω/dt gr dω/dt bd dω/dt nd dω/dt f(r) figure 1: ω̇gr vs ω̇gr,bd,nd,f(r) for the different relativistic theories (gr,bd,nd,f(r)): ω̇gr = ω̇gr, ω̇bd ∼= 0.92ω̇gr, ω̇nd ∼= 0.90ω̇gr, ω̇f(r) ∼= 1.10ω̇gr fig. 3 shows the isc’s comparison: log(k̄2obst h ) vs log(k̄2stellarmodel) are shown for different relativistic terms. the blue line is the trend of log(k̄2stellarmodel) = log(k̄2obst h ). it is evident that, apart from a few systems, log(k̄2stellarmodel) is always greater than the observed one. so the stellar core densities derived from the observations is higher than those coming from stellar model prevision. −0.02 0 0.02 0.04 0.06 0.08 0.1 0.12 0 0.5 1 1.5 2 2.5 3 3.5 x 10 −3 dω obs /dt d ω re l/d t dω rel /dt=dω obs /dt figure 2: observed ω̇obs vs ω̇rel for different relativistic terms (th ≡ gr,bd,nd,f(r)). the red line is the trend of ω̇rel = ω̇obs. it is evident that apart from some systems the relativistic term is always less than the observed one. 4 discussion and conclusion using data coming from apsidal motion rate of eedb we compared the variation of the relativistic term of the apsidal motion rate due to different theories of gravitation, that accordingly produces variation of classical newtonian term (see eq. 5). the results of this comparison was that we could not find any significant difference due to the effect of the different theories under test with respect to the standard general relativity. since the advance at periastron accumulates, the trend and the amount of this motion can be better determined by the observation of more orbits (or a larger fraction of orbit). longer observations also improve the determination of isc’s; thus, it would be possible, perhaps, to observe more significant differences, by checking the period variation on a much larger lapse of time and verifying the assumptions of syncronous orbital and rotation motion of the binary star components. a lot of observing work are producing new and more accurate data and step by step it is getting a better agreement between theory and observations. −4.5 −4 −3.5 −3 −2.5 −2 −1.5 −1 −2.6 −2.4 −2.2 −2 −1.8 −1.6 log k2obs th lo g k 2 s te ll ar m od el hp aur bm mon hh car v1765 cyg δ ori ι ori log k2 obs th =log k2 stellar model figure 3: internal second order structure constants (isc) log(k̄2obst h ) vs log(k̄2stellarmodel) for different relativistic terms (th ≡ gr,bd,nd,f(r)). the blue line is the trend of log(k̄2stellarmodel) = log(k̄2obst h ). it is evident that apart from some systems log(k̄2stellarmodel) is always greater than the observed one. so the stellar core according to the observations is more dense than the stellar model prevision. acknowledgement this work was supported by infn grant 2011.the authors gratefully acknowledge the referee prof. todor stanev for his useful suggestions. references [1] breen b., 1973, j. phys. a: math., nucl. gen., 6, 150. 257 leopoldo milano et al. [2] l. lin-sen, 2010, astrophys. space sci., 327, 59. doi:10.1007/s10509-010-0267-4 [3] giménez a., claret a., 2010, astron. and astroph. 519, a57. [4] levi civita t., 1937, amer. j. of math., 59, 225. doi:10.2307/2371404 [5] giménez a., 1985, astroph. j., 297, 405. doi:10.1086/163539 [6] thorne k.s., will c.m., 1971, astrophys. j., 163, 595. doi:10.1086/150803 [7] nordtvedt k., 1969, phys. rev., 180, 1293. doi:10.1103/physrev.180.1293 [8] biermann, p., et al. 1985, astrophys. j., 293, 303 [9] schiff l.i. , 1967, relativity theory and astrophysics vol 1, ed j ehlers 1967 (philadelphia: american mathematical society) [10] m. de laurentis,r. de rosa,f. garufi & l. milano,2012,mnras, 424, 2371-2379 [11] bulut i., demircan o., 2007, mon. not. r. astron. soc. 378, 179. [12] kopal z., dynamics of close binary systems. reidel,dordrecht (1978) [13] claret a., giménez a., 1992, astron. and astroph. suppl. 96, 255. [14] claret, a. 2004, a&a, 424, 919 doi:10.1111/j.1365-2966.2007.11756.x discussion jim beall: what systems would give a proper test? leopoldo milano: the systems that give a proper test are about eleven. carlotta pittori: do you think that it could be useful to study in more datail the outliers? which kind of obsevations could help to constrain the theory? leopoldo milano: many group of astronomers are doing new observations on the outliers with the aim of improving the knowledge on the rotations velocities and other parameters that can be the cause of the failure of the simple model that is generally adopted. 258 http://dx.doi.org/10.1007/s10509-010-0267-4 http://dx.doi.org/10.2307/2371404 http://dx.doi.org/10.1086/163539 http://dx.doi.org/10.1086/150803 http://dx.doi.org/10.1103/physrev.180.1293 http://dx.doi.org/10.1111/j.1365-2966.2007.11756.x introduction advance at periastron comparison with experimental data discussion and conclusion 227 acta polytechnica ctu proceedings 1(1): 227–230, 2014 227 doi: 10.14311/app.2014.01.0227 the galactic center region imaged by veritas from 2010–2012 matthias beilicke1 for the veritas collaboration2 1department of physics and mcdonnell center for the space sciences, washington university, st. louis, mo, usa 2http://veritas.sao.arizona.edu/ corresponding author: beilicke@physics.wustl.edu abstract the galactic center has long been a region of interest for high-energy and very-high-energy observations. many potential sources of gev/tev γ-ray emission are located in this region, e.g. the accretion of matter onto the central black hole, cosmic rays from a nearby shell-type supernova remnant, or the annihilation of dark matter. the galactic center has been detected at mev/gev energies by egret and recently by fermi/lat. at tev energies, the galactic center was detected at the level of 4 standard deviations with the whipple 10 m telescope and with one order of magnitude better sensitivity by h.e.s.s. and magic. we present the results from 3 years of veritas galactic center observations conducted at large zenith angles. the results are compared to astrophysical models. keywords: gamma-rays galactic center black hole non-thermal veritas. 1 introduction the center of our galaxy harbors a 4×106 m� black hole (bh) coinciding with the strong radio source sgr a*. x-ray/mev/gev transients in this region are observed on a regular basis. various astrophysical sources located close to the galactic center (gc) may potentially be capable of accelerating particles to multi-tev energies, such as the supernova remnant sgr a east or a pulsar wind nebula [1]. furthermore, super-symmetric neutralinos χ are discussed as potential candidates for dark matter accumulating in the gc region and annihilating into γ-rays [2]. the resulting spectrum would have a cut-off near the neutralino mass mχ. assuming a certain dark matter density profile the expected γ-ray flux along the line-of-sight integral can be calculated as a function of mχ and the annihilation cross section [3] and can in turn be compared to measured upper limits. egret detected a mev/gev source 3eg j17462851 coincident with the gc position [4] and recently fermi/lat resolved several sources in the gc region [5], see fig. 3. however, uncertainties in the diffuse galactic background models and limited angular resolution at mev/gev energies make it difficult to study the morphologies of these sources. at gev/tev energies a detection from the direction of the gc was first reported in 2001/02 by the cangaroo ii collaboration with a steep energy spectrum dn/de ∝ e−4.6 at the level of 10% of the crab nebula flux [6]. shortly after, evidence at the level of 3.7 standard deviations (s.d.) was reported from the whipple 10 m collaboration [7]. the gc was finally confirmed as a gev/tev γ-ray source by the h.e.s.s. collaboration [8] (the position of the supernova remnant sgr a east could be excluded as the source of the γ-ray emission). the energy spectrum measured by h.e.s.s. is well described by a power-law dn/de ∝ e−2.1 with a cut-off at ∼15 tev. the h.e.s.s. observations revealed a diffuse gev/tev γ-ray component (dashed contour lines in fig. 3) which is aligned along the galactic plane and follows the structure of molecular clouds [9]; the emission is explained by an interaction of local cosmic rays (crs) with matter of the molecular clouds. the magic collaboration detected the gc (7 s. d.) in 2004/05 observations performed at large zenith angles (lza) [10], followed by a strong (> 10 standard deviations) veritas lza detection in 2010 [11]. 2 veritas observations of the galactic center gc observations due to its declination the gc can only be observed by veritas at lza (zenith angles 60 − 66 deg) – strongly decreasing the angular resolution and sensitivity. the use of the displacement parameter [12], between the center of gravity of the image and the shower position, has been used in the veritas event reconstruction which strongly improved the sensitivity for lza observations [11]. the performance and energy reconstruction have been confirmed on lza crab nebula data. the column density of the atmosphere changes with 1/ cos(z). in a conservative es227 http://dx.doi.org/10.14311/app.2014.01.0227 matthias beilicke timate, the systematic error in the energy/flux reconstruction can be expected to scale accordingly. more detailed studies are needed for an accurate estimate; for the gc observations we currently give a conservative value of a systematic error on the lza flux normalization of ∆φ/φ ' 0.4. the gc was observed by veritas in 2010–2012 for 46 hrs (good quality data, dead-time corrected) with an average energy threshold of ethr ' 2.5 tev. gc results the veritas sky map of the gc region is shown in fig. 3. an 18 s.d. excess is detected. no evidence for variability was found in the 3-year data. the energy spectrum is shown in fig. 1 and is found to be compatible with the spectra measured by whipple, h.e.s.s., and magic. since the large lza effective areas of the veritas observations compensate a shorter exposure of low-zenith observations, the statistical errors of the e > 2.5 tev data points are comparable or even smaller than those of the h.e.s.s. measurements. ] -1 s -2 d n /d e [ e rg c m 2 e -1310 -1210 -1110 energy [ev] 1210 13 10 v e r it a s ( 2 0 1 0 -2 0 1 2 , p re li m .) galactic center veritas (gc) h.e.s.s. (gc) whipple (gc) magic (gc) figure 1: veritas energy spectrum measured from the direction of the gc (statistical errors only). also shown are bow ties representing the spectra measured by whipple [7], h.e.s.s. [8], and magic [10]. diffuse flux limit and dark-matter annihilation off-source observations were performed in a field located in the vicinity of the gc region (similar zenith angles and sky brightness) without a known tev γ-ray source. these observations are used to study the background acceptance throughout the field of view and allow the estimate of a diffuse γ-ray component surrounding the position of the gc. an upper limit of the diffuse γ-ray flux can in turn be compared with line-of-sight integrals along the density profile ∫ ρ2dl, in order to constrain the annihilation cross section for a particular dark matter model, dark matter particle mass and density profile ρ(r). due to its likely astrophysical origin the excess at the gc itself, as well as a region along the galactic plane, will be excluded from this analysis (work in progress). hadronic models hadronic acceleration models [13, 14] involve: (i) hadrons being accelerated in the bh vicinity (few tens of schwarzschild radii). (ii) the accelerated protons diffuse out into the interstellar medium where they (iii) produce neutral pions which decay into gev/tev γ-rays. linden et al. (2012) discuss the surrounding gas as proton target defining the morphology of the tev γ-ray emission [15]. changes in γ-ray flux in those models can be caused by changing conditions in the bh vicinity (e.g. accretion). the time scales of flux variations are ∼104 yr at mev/gev energies (old flares) and ∼10 yr at e > 10 tev (’new’ flares caused by recently injected high-energy particles) [13]. constraining the e > 10 tev spectral variability would serve as an important test for this class of models. ] -1 s -2 d n /d e [ e rg c m 2 e -1310 -1210 -1110 -1010 energy [ev] 8 10 1110 1410 veritas (preliminary) galactic center veritas (gc) fermi (chernyakova et al., 2011) ballantyne et al. (2011) [hdr] chernyakova et al. (2011) [hdr] linden et al. (2012) [hdr] atoyan et al. (2004) [lep] figure 2: veritas energy spectrum compared to hadronic [13, 14, 15] and leptonic [16] emission models discussed for the gc source. the fermi/lat bow tie is taken from [13]. leptonic models atoyan et al. (2004) [16] discuss a bh plerion model in which a termination shock of a leptonic wind accelerates leptons to relativistic energies which in turn produce tev γ-rays via inverse compton scattering. the flux variability time scale in this model is on the order of tvar ∼100 yr. the hadronic and the leptonic models are shown together with the veritas/fermi data in fig. 2. the leptonic model clearly fails in explaining the flux in the mev/gev regime. however, this emission may well originate from a spatially different region or mechanism other than the tev γ-ray emission. the hadronic models can explain the sed by the superposition of different flare stages. future fermi/veritas flux correlation studies, as well as the measurement of the tev energy cut-off and limits on the e > 10 tev variability will serve as crucial inputs for the modeling. 3 summary and conclusion veritas is capable of detecting the gc within 3 hrs in observations conducted at zenith angles greater than 228 the galactic center region imaged by veritas from 2010–2012 figure 3: veritas sky map of the gc region (smoothed excess significances, ring background, scale truncated). the black contour lines indicate the gc and the supernova remnant g 0.9+0.1 as seen by h.e.s.s. [8]. the gray dashed lines indicate the h.e.s.s. diffuse emission along the galactic plane and from hess j1745-303 [9]. the position of hess j1741-302 is indicated, as well (circle); the flux/spectrum of this source make it very unlikely to be detected in veritas lza observations. the solid circles (cyan color) indicate the positions of the mev/gev sources taken from the second fermi/lat catalog [5]. 60 deg. the measured energy spectrum is found to be in agreement with earlier measurements by h.e.s.s., magic, and whipple. future observations to measure the cut-off energy in the spectrum and to determine limits on the flux variability at the highest energies will place constraints on emission models. the recently discovered giant molecular cloud heading towards the immediate vicinity of the gc bh [17] represents further motivation for future tev γ-ray monitoring of this region. an upper limit on diffuse γ-ray emission and, in consequence, a limit on the photon flux initiated by the annihilation of dark matter particles is work in progress. acknowledgement this research is supported by grants from the u.s. department of energy office of science, the u.s. national science foundation and the smithsonian institution, by nserc in canada, by science foundation ireland (sfi 10/rfp/ast2748) and by stfc in the u.k. we acknowledge the excellent work of the technical support staff at the fred lawrence whipple observatory and at the collaborating institutions in the construction and operation of the instrument. references [1] q.d. wang, f.j. lu, e.v. gotthelf, et al., mnras 367, 937 (2006). doi:10.1111/j.1365-2966.2006.09998.x [2] g. jungman, m. kamionkowski, & k. griest, phr 267, 195 (1996). [3] l. bergström, p. ullio, & j. buckley, aph 9, 137 (1998). [4] r.c. hartman, d.l. bertsch, s.d. bloom, et al., apjs 123, 79 (1999). [5] a.a. abdo, et al., apjs 188, 405 (2010). doi:10.1088/0067-0049/188/2/405 [6] k. tsuchiya, r. enomoto, l.t. ksenofontov, et al., apj 606, l115 (2004). doi:10.1086/421292 [7] k. kosack, h.m. badran, i.h. bond, et al., apj 608, 97 (2004). doi:10.1086/422469 [8] f.a. aharonian, et al., a&a 425, l13 (2004). [9] f.a. aharonian, et al., nature 439, 695 (2006). doi:10.1038/nature04467 [10] j. albert, et al., apj 638, l101 (2006). doi:10.1086/501164 [11] m. beilicke, et al. (veritas collaboration), proc. 2011 fermi symp., arxiv 1109.6836 (2011). [12] j.h. buckley, c.w. akerlof, d.a. carter-lewis, et al., a&a 329, 639 (1998). [13] m. chernyakova, d. malyshev, f.a. aharonian, et al., apj 726, 60 (2011). doi:10.1088/0004-637x/726/2/60 229 http://dx.doi.org/10.1111/j.1365-2966.2006.09998.x http://dx.doi.org/10.1088/0067-0049/188/2/405 http://dx.doi.org/10.1086/421292 http://dx.doi.org/10.1086/422469 http://dx.doi.org/10.1038/nature04467 http://dx.doi.org/10.1086/501164 http://dx.doi.org/10.1088/0004-637x/726/2/60 matthias beilicke [14] d.r. ballantyne, m. schumann, & b. ford, mnras 410, 1521 (2011). [15] t. linden, e. lovegrove, & s. profumo, arxiv 1203.3539 (2012). [16] a. atoyan, & c.d. dermer, apj 617, l123 (2004). doi:10.1086/427390 [17] s. gillessen, r. genzel, t. fritz, et al., nature 481, 51 (2012). doi:10.1038/nature10652 230 http://dx.doi.org/10.1086/427390 http://dx.doi.org/10.1038/nature10652 introduction veritas observations of the galactic center summary and conclusion 278 acta polytechnica ctu proceedings 1(1): 278–282, 2014 278 doi: 10.14311/app.2014.01.0278 havms: highly available virtual machine computer system fault tolerant with automatic failback and close to zero downtime memmo federici1, carlo gaibisso2, bruno l. martino3 1istituto di astrofisica e planetologia spaziali, inaf iaps via fosso del cavaliere 100, 00133 roma, italy 2istituto di analisi dei sistemi ed informatica ”antonio ruberti”, iasi-cnr viale manzoni 30, 00185 roma , italy 3associated inaf iaps corresponding author: memmo.federici@iaps.inaf.it abstract in scientific computing, systems often manage computations that require continuous acquisition of of satellite data and the management of large databases, as well as the execution of analysis software and simulation models (e.g. monte carlo or molecular dynamics cell simulations) which may require several weeks of continuous run. these systems, consequently, should ensure the continuity of operation even in case of serious faults. havms (high availability virtual machine system) is a highly available, ”fault tolerant” system with zero downtime in case of fault. it is based on the use of virtual machines and implemented by two servers with similar characteristics. havms, thanks to the developed software solutions, is unique in its kind since it automatically failbacks once faults have been fixed. the system has been designed to be used both with professional or inexpensive hardware and supports the simultaneous execution of multiple services such as: web, mail, computing and administrative services, uninterrupted computing, data base management. finally the system is cost effective adopting exclusively open source solutions, is easily manageable and for general use. keywords: havms high availability fault tolerant open source multitask. 1 introduction havms effectively solves the problems related to the ”robustness” of computer systems wholly embracing the concept of high availability. a system in high availability, ha in what follows, must ensure the continuity over time of the provided services, which, in case of fault, must be restored in the shortest possible time. havms, through an accurate design of the hw and sw, significantly reduces the faults and their negative effects on the provide services. havms has been designed mainly keeping in mind the following requirements: cost effectiveness, ease of management and, above all, the ability to automatically implement restoring strategies of the provided services without any interruption. all the above requirements have been met by an accurate choice of the available open source solutions meeting our targets and their integration with the sw specifically conceived by the authors. in particular, the last of them makes the system unique in its kind and competitive with analogous commercial solutions. havms is made by two servers, one active and the other dormant, but from the point of view of users, as well as from that of applications, the system is seen as a single server. the fault tolerance of the system is ensured by a continuous synchronization of the two servers. this synchronization keep the data and the states of all the virtual machines, vms in what follows, perfectly aligned. every service runs on a different vm each of which has its own ram, storage and some computing resources. each vm is independent of the others. when a fault occurs, the sleeper server is awakened and automatically takes the place of the broken server. once the fault is fixed, it will be enough to reconnect the repaired server and automatically the failback procedure restores the proper functioning of the system including its ha capabilities. 1.1 the context havms is a general purpose system, this is one of its strengths. among the potential areas in which the system proved its effectiveness, there is the acquisition and processing of data from space missions. the main abilities required to havms by this particular context are: • managing of processes that require uninterrupted data analysis and data acquisition; 278 http://dx.doi.org/10.14311/app.2014.01.0278 havms: highly available virtual machine computer system fault... • storing the results of the analysis for long periods of time; • making these results available to the scientific community, in an effective, continuos and reliable way. in fact, scientists in this area are interested in performing their analysis and tests on spatial observations as soon and on as much data as possible. furthermore, missions such as integral (international gamma-ray astrophysics laboratory) winkler, et al. (2003), have to maintain for a long period of time the results of the scientific analysis performed on the whole data set and usually released, as surveys, about every year. finally a further important requirement that has been met is making the management of the whole system, as far as possible, easy and automatic, thus minimizing the costs related to human resources devoted to the management itself. 1.2 the alternatives in designing, havms we deeply analyze and consider as points of reference the more widespread alternatives currently implementing the concept of ha. all these alternatives, in general, suffers from the problem of the adoption of often expensive sw and hw architectures. more in details these are, from our point of view and with respect to the particular operation context we are considering, the main drawbacks of the considered solutions: • windows server failover clustering (wsfc): requires the purchase of a quite expensive sw license (when this paper is written, about 900 euros per processor). all servers in the cluster must be absolutely identical. ha is guaranteed by a mirroring system, replicating the status of the severs every 5-10 minutes: if there is a fault during a replica, the replica itself may fail and the running jobs are aborted with a high probability. • vmware vsphere: requires the purchase of a quite expensive enterprise license (when this paper is written, about 4000 $). requires professional hw and an additional server to handle the nodes in the system. the setup is not simple. • red hat cluster manager (red hat enterprise linux server): requires the purchase of an annual license (when this paper is written, about 500 $). in case of fault the system must be restarted, as a consequence all runs in progress are stopped. does not support an automatic failback mechanism. 2 our solution figure 1: havms block diagram fig.1 shows the general architecture of havms. continuity of service is provided through an automatic failover (activates the dormant server in case of fault of the active one) and failback (restores the initial system configuration and operation) mechanisms. the software components implementing these mechanisms are: • xen vms manager, a manager of virtual machines; • drbd (protocol d), a synchronizer of block devices; • remus, provides transparent, comprehensive high availability to ordinary vms running on xen; • some components, designed and developed by the authors especially conceived, among the others, to implement the automatic failback and system startup. the system is based on the services provided by xen vms manager. each vm has its own ip address by which it is accessible from the outside. as previously stated, one server is active, the other one is dormant. the two servers are connected through a lan at 1 gb/sec. remus (in green) and drbd (in red) seamlessly (once every 40 msec) replicates the state of the active server on the dormant one. in particular, remus replicates the states of vms and simultaneously sends a trigger to drbd, which, in turns, synchronizes 279 memmo federici, carlo gaibisso, bruno l. martino the block devices (hard disks) of the two servers. as a consequence the two servers are aligned once every 40 msec. users connect to the vms through their unique ip address. vms makes it possible to install different operating systems, on which different softwares and services can be run, such as: matlab, idl and web servers, compilers and so on. 3 servers software and hardware architecture both servers have the same software architecture, already introduced in the previous section. the whole architecture leans on linux ubuntu server 10.04 64 (with a kernel customized to support xen). in what follows the single components of the architecture are briefly described. 3.1 xen xen hypervisor is a virtualization platform licensed under the gpl developed at the computer laboratory of the university of cambridge. xen is included in all major linux distributions and increasingly adopted by commercial solutions. one of the its most interesting feature is the ability to effectively control the requests of access to physical resources coming from vms through a paravirtualization mechanism. this mechanism guarantees a minimum decay of performance due to virtualization, since requests of access coming from vms are mainly executed on the physical computing resources. paravirtualization requires a customized version of the ubuntu kernel. 3.2 drbd drbd (distributed replicated block device) is a distributed storage system for the gnu/linux platform, usually adopted by ha clusters. drbd is responsible for the synchronization of data between the servers: one of them is identified as primary, the other one as secondary. when the primary server fails, a management process promotes the secondary one to the role of primary. when the fault is fixed, the system may reestablish the roles initially assigned to the servers, after a resynchronization of the data storage devices. this synchronization is particularly effective since only those blocks that were changed during the outage are resynchronized. 3.3 remus remus is part of xen and implements the ha concept by replicating on the dormant server the state of all the vms machine running on the active server. this snapshot occurs once every 40 msec (this value can be changed at setup time); at the same time remus sends a trigger to drbd for the synchronization of the hard disks. in this way the active server is constantly aligned, both for what concerns data and computational aspects, with the dormant server. in case of fault of the active server, the dormant one is awakened and becomes immediately active, thus avoiding any interruption in the provisioning of services. moreover, the tcp/ip protocol guarantees the correct transmission of data packets, even when the connectivity is temporary interrupted. the architecture required to effectively support our ha solution, does not expect the use of professional hw. in fact, as will be evident from the following description of our experiences in the field, an entry level solution with the following characteristics turns out to be absolutely appropriate to the achievement of our goals: • i7 8 cores intel processor • 8/16 gb ram ddr3 • 2 tb hard disk • 2 network interfaces 3.4 failover the typical faults of a operating environment such as the one here considered here, can be attributed mainly to two categories: lack of connection, malfunctioning of some vms. in case of a fault, the failover process is automatically triggered, which: 1. stops the drbd synchronization between servers; 2. if running, stops the execution of remus on the active server; 3. awakens the dormant server. the execution of vms is consequently resumed from the last committed checkpoint (at most 40 msec before). 3.5 failback unfortunately, remus does not support any kind of automatic failback. the sw developed for havms compensates for this deficiency and gives it its unicity. this process, which is automatically triggered and implemented, restores the initial configuration and operational capabilities of the system once the fault has been fixed, included its ha functionality. 280 havms: highly available virtual machine computer system fault... 4 practical applications the development of havms was determined by some practical needs. below we briefly describe the two applications that currently rely on the services offered by the system. 4.1 the ha data storage system of integral the laboratory of distributed computing at iaps (inaf) is in charge of aves federici et al. (2012), the cluster devoted to the analysis of data collected by integral. aves is connected to data storage subsystem (dss) a dedicated storage system martino and federici (2011) adopting havms. dss automatically downloads from integral science data center (isdc) courvoisier et al. (2003) the data collected by the satellite, backups this data and makes them available to aves (16 tb to rise). as a consequence of a fault of dss the download is interrupted and aves as no longer access to the shared data. this might have severe consequences on the ongoing activities since, in turn, may cause a data misalignment and a crash of the running analysis. havms avoid these risks and their unwanted consequences. 4.2 continuity of service at iasi havms guarantees the continuity of the service provided by iasi it division, among them: attendance control, centralized computing, storage and backup, printing services management. a different vm is allocated to each service. 5 further potential applications in the following we describe two potential fields of application of havms we are investigating. these fields differ substantially from one another. each of them is representative of a significant and large family of applications. 5.1 high end solution large structures, such as hospitals or national administrative offices should provide a relevant number of services that are usually hosted on individual servers. these servers, to ensure an acceptable level of reliability, should be equipped by systems for the backup of the status of the provided services and the data they deal with. this is therefore a privileged context within which to exploit the characteristics of havms. a quick market survey has shown that our solution can meet the above mentioned reliability requirements with not negligible costs but still low if compared with those of similar commercial solutions. in fact, one of the adoptable hw solutions could be the following: motherboard supermicro xeon mp series x9qxxx, equipped with 4-socket xeon processors (32 cores), 120 gb of ram, 4 ethernet interfaces etc. this configuration may provide up to 30 services, each with its own vms and ip address, with a cost of approximately 5000 euros per server. 5.2 a little gem in this section we describe a solution adopting havms that fundamentally, both in terms of computing power and cost, differs from the high end solution, which is a good example of the great versatility of our ha solution. this solution has been designed to deal with the monitoring of atmospherical and geological events by the use of sensors in hard to reach sites. the resulting system, which has a cost of about 300$ when this paper is written, is composed by three raspberry pi collecting analog and digital signals from sensors and transferring them by a radio link to a remote control station. thanks to the low power consumption, the system can be powered by a small photovoltaic panels. two of the three raspberry’s are allocated to havms; the third one, which is initially turned off, is switched on in case of failure of any one of the first two. this is obtained by a hw signal sent through lan (wakeup on lan). the awakened raspberry takes the place of the broken one thus guaranteeing the continuity of service, included those offered by havms. this solution substantially limits the number of maintenance interventions, thereby reducing the relative costs. figure 2: the raspberry pi three node havms 6 further work we plan to improve the ha capabilities of havms by adding to the current configuration a third server in 281 memmo federici, carlo gaibisso, bruno l. martino wake on lan. in case of fault this third server is automatically awakened by an hw signal sent through the lan and takes the place of the broken one. this automatic mechanism should not require any intervention by the system manager. with this solution, the system maintains its characteristics of ha also in case of fault of one server. to improve the effectiveness of havms in relation to the use of network resources, multiple network connections can be combined in parallel (bonding) to increase throughput beyond what a single connection could sustain, and to provide redundancy in case one of the links fails. 7 conclusions and discussion havms is a highly available fault tolerant general purpose, recyclable system based on the use of vms, assuring continuity of operation and no interruption in services providing in case of fault. havms is cost effective since it only adopts open source solutions. it is also extremely versatile, providing all operating systems supported by xen. paravirtualization dramatically reduces the negative impact of virtualization on the performance of the whole system. the system supports automatic failover and failback within times close to zero. automatic failback is an exclusive feature of havms. acknowledgement the authors thanks to: giuliano sabatino, fabio guglietta. references [1] winkler, c., et al.: the integral mission. astron. astrophys. 411, l1l6 (2003) [2] m. federici, et al. 2012, aves: a high performance computer cluster array. exp astron (2012) 34:105121 [3] b.l. martino and m. federici ”an high availability data storage subsystem for the integral data analysis”, mem. s.a.it. vol. 83, 377 2011 [4] courvoisier, t.j.-l., et al.: the integral science data centre (isdc) for the integral satellite scientific data analysis. astron. astrophys. 411, l53l57 (2003) doi:10.1051/0004-6361:20031172 discussion beall james’s question: do you plan to use the raspberry pi machine for computation? memmo federici’s answer: not in the traditional sense because rb is equipped with small amount of ram (only 256 mb) and little computing power. there are some applications that see this hw used in small clusters with low consumption. we plan to use rb to make some services such as web servers and print servers. this small device is very versatile and lends itself also to the process control. – 282 http://dx.doi.org/10.1051/0004-6361:20031172 introduction the context the alternatives our solution servers software and hardware architecture xen drbd remus failover failback practical applications the ha data storage system of integral continuity of service at iasi further potential applications high end solution a little gem further work conclusions and discussion acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0022 acta polytechnica ctu proceedings 5:22–25, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app infrastructure parameters affecting capacity of railways in ten-t vít janoš∗, milan kříž department of logistics and management of transport, faculty of transportation sciences, ctu in prague, czech republic ∗ corresponding author: janos@fd.cvut.cz abstract. the article presents possible solutions of the issue of the lost capacity on a double-track line. it is shown a solution where an inserted train on the double-track line use the lost capacity of the opposite line track using active overtaking. the feasibility of such a solution is discussed and the length of sections suitable for active overtaking is calculated. keywords: capacity, uic 406, speed ratio . 1. introduction the capacity of the railway infrastructure can be considered as a usable time space for the train paths. capacity is thus influenced not only by the parameters of the infrastructure, but also by the character of the traffic – by the designed train paths and their heterogeneity. with increasing degree of heterogeneity the capacity consumption increases too and the number of possible train paths in a defined time windows declines. this phenomenon is also called "capacity balance" [1]. in the capacity balance there are four factors on whose interdependencies the capacity is based: number of trains, average speed, stability and heterogeneity. relationships among the individual factors are shown in figure 2 [1]. principle of figure 2 is explained as follows: "in this qualitative model, an axis for each parameter is drawn from a unique origin. a chord links the points on the axes, corresponding to the value of each parameter. the length of the chord represents the capacity. capacity utilisation is defined by the positions of the chord on the four axes. increasing capacity means increasing the length of the chord". general effect associated with the modernization of the trans-european conventional rail network (typically the modernization of the transit railway corridors in the czech republic in the original routes to a maximum speed of 160 km/h) is the rising heterogeneity of the traffic on some routes in the network. it is caused by increasing the line speeds. on the one hand regional passenger train paths remains after increasing the line speed about as fast as before the modernization (journey speed of the regional trains usually does not exceed 60 km/h due to the higher number of stops). on the other hand long-distance passenger trains should logically utilize all characteristics of infrastructure for the maximum journey speed (at the fastest trains the journey speed of 120-140 km/h can be reached in some line sections). these both cases are thus usually marginal and between these limits figure 1. possible situation of active overtaking on a double-track line. the freight trains paths occur (with journey speeds in the range of 70-100 km/h). 2. heterogeneity and its measuring heterogeneity can be measured in different ways. one way is to compare running speeds of individual train (or train paths). for this purpose krueger [2] proposed and used so called speed ratio (sr). the speed ratio is the ratio of the fastest train speed to the slowest train speed: sr = max(v1, v2, ..., vn) min(v1, v2, ..., vn) so the sr on the typical line section on some transit railway corridor in the czech republic can reached value of 3 (the speed of the fastest train about 140 km/h and the speed of the slowest train about 45-50 km/h). 22 http://dx.doi.org/10.14311/app.2016.5.0022 http://ojs.cvut.cz/ojs/index.php/app vol. 5/2016 infrastructure parameters affecting capacity of railways in ten-t figure 2. capacity balance [1]. the very high sr is problematic only when it comes to the line or line section with a high number of train paths. but the lines of the trans-european conventional rail network usually show a high number of train paths. line sections with high demand for train paths and a high sr can usually be solved by segregating the various types of traffic. this entails increasing the number of line tracks (e.g. quadruple-track line sections in urban agglomerations with two tracks for long-distance passenger trains and fast freight trains and two tracks for slower regional passenger trains and slower freight trains) or construction of lines dedicated exclusively for a certain type of traffic (typically on the one hand high-speed railway lines and, on the other hand, lines dedicated only for freight train as e.g. betuwelijn). this approach, however, is not yet common in the czech republic. 3. active overtaking on double-track lines, where high degree of heterogeneity is because of the mixed traffic and high requirements for the number of the train paths, so called lost capacity arises. lost capacity is unusable time space between train paths with different train speeds, where any additional train path cannot be inserted (e.g. because there is no other station). lost capacity can be limited by reducing heterogeneity of train paths. either the slowest trains can be accelerated (e.g. these trains can pass some stops), or vice versa the fastest trains can be slowed. lost capacity may be also used for active overtaking1. an example of this solution is shown in the figure 1. active overtaking is a very interesting possibility of using a part of the lost capacity. this idea can be very helpful in cases, when the number of stations on a line section is reducing within modernization of infrastructure. in the former stations there are only crossovers for crossing between the line tracks for cases of interruptions. in such cases there is no possibility to insert one more train path into a sequence of differently 1active overtaking can be defined as an overtaking of slower train by faster train, when the slower train does not have to stop [3]. fast train paths, but they can also be used for crossing between line tracks and in regular operation. the actual possibility of using these crossovers for regular traffic, however, depends on the diverging speed that affects the train speed (and thus journey time) of the overtaking or the overtaken train and thus affects the length of the section required for overtaking. furthermore, it is obvious that the greater the heterogeneity of the train paths in the reference section is (i.e. the larger the difference of the journey speeds between the overtaken and the overtaking train), the shorter the section required for overtaking is. with regard to the above mentioned borderline cases it is evident that the overtaken slower train is always a regional passenger train. there are generally two cases possible: (1.) the slower train crosses over to the other line track and runs against the right direction, the faster train runs in the right direction. (2.) the faster train crosses over to the other line track and runs against the right direction, the slower train runs in the right direction. the cases from the first group are suitable for situations, in which the platforms of stations and stops in the line section are situated between the line tracks, or in which only stations equipped with a good visual and audible information system are. this system should provide clear information about platforms and tracks of arriving and departing trains. the cases from the second group are suitable, if the speed limit over the crossovers between the line tracks corresponds or is close to the line speed or the maximum train speed. 4. case study: the first transit railway corridor between prague and děčín the part of the first transit railway corridor between prague and děčín links the prague metropolitan area [4] with the ústí metropolitan area and with the north part of germany. the most typical situation for the active overtaking is an interaction of a fast freight train (maximum train speed about 100 km/h) and a regional passenger train (journey speed about 50 km/h). the active overtaking of the first type can be used in the line sections: • hněvice – hněvice seř.n. • roudnice nad labem – hrobce • lovosice jih – lovosice • prackovice nad labem – ústí nad labem jih – ústí nad labem hl.n. now we have to analyse these line section, whether the active overtaking leads to time saving in comparison with overtaking of the regional passenger train 23 vít janoš, milan kříž acta polytechnica ctu proceedings figure 3. relationship between speed ration and length of section for overtaking. by the fast freight train, when the regional passenger train is dwelling in some station. sections hněvice – hněvice seř.n. and lovosice jih – lovosice can be excluded immediately, because the result time savings will be negative. in both cases the regional passenger train is stopping only once in the line section and in both cases the speed limit over crossovers in the stations which bounds the line section is lower than the speed limit on passing sidings in these stations. sections roudnice nad labem – hrobce and prackovice nad labem – ústí nad labem jih – ústí nad labem hl.n. show always positive time savings. in both cases the regional passenger train is stopping twice in the line section2. in these cases, the parallel run is generally about 2 minutes more effective then overtaking of the slower train in the station. it is very important that the overtaking train is guided precisely to the desired time position. the accuracy of this guidance may have a significant effect on the resulting time savings. by the active overtaking of the second type the journey time of the overtaking train is extended because of the diverging speed of the crossovers in the bounding stations. in table 1 values for typical situation on the first transit railway corridor are shown. the values were calculated with linear margin 9%. value are graphically illustrated in figure 3. from table 1 it is possible to determine the length of the line section which is needed for active overtaking. if the actual line section is shorter, then the dwell times of the slower train have to be extended. 5. conclusions the verification of heterogeneity in relation to the partial use of the lost capacity by active overtaking of differently fast trains showed clearly that in the case of the interaction of regional passenger trains and freight trains there are solutions which can be divided into 2also the stops in bounding stations of the line section are counted, if the regional passenger is being already overtaken during the stay in the station. sum of operating intervals and time reserves [min] 5 journey speed of slower train [km/h] 50 speed ratio [1] speed of faster train [km/h] length of section for overtaking [km] 1,1 55 45,83 1,2 60 25,00 1,3 65 18,06 1,4 70 14,58 1,5 75 12,50 1,6 80 11,11 1,7 85 10,12 1,8 90 9,38 1,9 95 8,80 2 100 8,33 2,1 105 7,95 2,2 110 7,64 2,3 115 7,37 2,4 120 7,14 2,5 125 6,94 2,6 130 6,77 2,7 135 6,62 2,8 140 6,48 2,9 145 6,36 3 150 6,25 table 1. length of section for overtaking. two situations. either if the diverging speed over the crossovers in the bounding station of the line section is clearly lower than the running speed of the fast freight train, then the regional passenger train should run against the right direction on the other line track. or if the diverging speed over the crossovers in the bounding station of the line section is high enough (minimum is 80 km/h), then then the fast freight train should run against the right direction on the other line track. it is important to mention that by modernizing the infrastructure, which should also allow such use of the lost capacity, platforms should be mostly located between the main line tracks. with regard to the actual length of the line sections for active overtaking and with regard to demand for the train paths it is necessary to mention that on 24 vol. 5/2016 infrastructure parameters affecting capacity of railways in ten-t trans-european conventional rail network this solution will be usually marginal. furthermore, it is likely that the heterogeneity of demanded train paths will not decline. authorities ordering the long distance passenger trains and open access operators do not want to, on the one hand, slow down artificially their trains. on the other hand, local authorities ordering the regional passenger trains do not want to reduce the number of operated railway stops, which is a very problematical political topic. in the czech republic the construction of quadrupletrack line sections remains as a permanent challenge in cases where a high degree of heterogeneity connected with high demand for train paths occurs. only such a solution leading to the segregation of various types of traffic allows homogenization of the train paths and effective elimination of the lost capacity. references [1] uic leaflet 406 capacity. paris: international union of railways (uic). [2] h. krueger. parametric modeling in rail capacity planning. simulation conference proceedings. az, phoenix, 1999. [3] m. drábek. periodic freight train path in network. čvut, praha, 2014. [4] t. kostelecký, d. čermák. metropolitan areas in the czech republic definitions, basic characteristic, patterns of suburbanisation and their impact on political behaviour. sociologický ústav av čr, 2004. 25 acta polytechnica ctu proceedings 5:22–25, 2016 1 introduction 2 heterogeneity and its measuring 3 active overtaking 4 case study: the first transit railway corridor between prague and decín 5 conclusions references 128 acta polytechnica ctu proceedings 2(1): 128–132, 2015 128 doi: 10.14311/app.2015.02.0128 investigating long-term behaviour of x-ray binaries using archival data m. m. kotze1,2, p. a. charles2,3 1south african astronomical observatory, p.o. box 9, observatory 7935, south africa (sa) 2astrophysics, cosmology and gravity centre (acgc), astronomy department, university of cape town, rondebosch 7701, sa 3school of physics & astronomy, university of southampton, southampton so17 1bj, uk corresponding author: marissa@saao.ac.za abstract long term modulations have been detected in a wide variety of both low and high-mass x-ray binaries. the all sky monitor on board the rossi x-ray timing explorer provides the most extensive (∼15 years) and sensitive x-ray archive for studying such behaviour. since those variations were often intermittent and/or aperiodic, we used a time-dependent dynamic power spectrum method to examine how the modulations themselves vary with time in a systematic way. some were found to be remarkably stable, while others show a range of properties, from even longer variability time-scales to quite chaotic behaviour. keywords: accretion discs x-ray binaries. 1 introduction kotze & charles (2012) contains the results of our timedependent period analysis on 25 x-ray binaries (xrbs) with reported long term variability. for the purpose of these conference proceedings, some of those results are reproduced herein to illustrate the importance of such an approach when dealing with varying periodic and/or aperiodic signals. we also use this opportunity to present some of the highlights contained in our previously published results. 2 data analysis 2.1 rxte/asm archival data the massachusetts institute of technology (mit) operated an all sky monitor (asm) on board the rossi x-ray timing explorer (rxte) from 1996 to 2012. the asm observed the x-ray sky with 3 rotating scanning shadow cameras (ssc) which scanned ∼80% of the sky during ∼90 min orbits. data were reduced and compiled weekly by the asm team and made publicly available 1 as dwell-by-dwell or one-day-averages in four energy bands, which includes a sum-band (1.5-12 kev). one-day-average data 2 are the dwell-by-dwell data binned into 1-day bins. archival asm datasets contain the lightcurves of all xray sources in the rxte catalogue. for full details see levine et al. (1996). 2.2 time-dependent period analysis periods are typically found by variability analysis of time-series/lightcurve datasets. while this approach may allow distinction between periodic and quasiperiodic behaviour, it contains no information regarding the variation or consistency of aperiodic/quasi-periodic signals. clarkson et al. (2003) presented results from their time-dependent periodic analysis on 4 x-ray binaries in the form of dynamic power spectra (dps). the method requires the datasets to be split into windows that are of sufficient length to allow detection of the maximum period considered. transition from one independent window to the next is smoothed (i.e. increased resolution) by adding sliding windows between the positions of adjacent independent data windows. lomb-scargle (l-s; lomb 1976 & scargle 1982) periodograms were produced for every data window and the results of all the periodograms for a source were plotted together in a density map, using the l-s power for each frequency plotted at every window’s mid-point along the time-axis. the frequency domain allowed appropri1http://xte.mit.edu/ and http://xte.mit.edu/asmlc.html 2i have made all their plots available on: http://www.saao.ac.za/∼marissa/lc/lc.html 128 http://dx.doi.org/10.14311/app.2015.02.0128 investigating long-term behaviour of x-ray binaries using archival data ate coverage over frequencies associated with previously reported periodic behaviour (table 1). since the dps method requires a significant number of computations, its application has been limited until recent advances in desktop technology allowed it to be employed more readily. figure 1: accretion disc stability to radiation-driven warping in xrbs, as functions of q and rb ( [ gm1 c2 ] ), adapted from ogilvie & dubus (2001) to include only xrbs with known super-orbital periods and updated with their latest system parameters (see table 1). squares indicate low-mass xrbs (lmxbs) and circles high-mass xrbs (hmxbs). tidally-induced disc precession (sec 3.4) may occur left of the dashed line. 3 results several mechanisms have been proposed to account for the super-orbital variations (psup) observed in xrbs. we list our results according to the mechanisms that have been considered most likely candidates for the cases included here. please refer to kotze & charles (2012) for more details. 3.1 radiation-induced disc warp/tilting ogilvie & dubus (2001) provided stability predictions for xrb accretion discs against radiation-driven warping/tilting. therein the binary separation (rb) and mass ratio (q = m2/m1) determine the location of a source on their predictive diagram, which is divided into zones where stable warps (between the dashed and solid line), chaotic warps (above the dashed line) or no warps (below lower solid line) are expected. a source near the transition boundary between the stable warp zone and the chaotic warp zone, may experience instability that could result in variable warps. table 1: system parameters for included xrbs source psup porb q rb/10 6 [days] [days] [ m2 m1 ] [ gm1 c2 ] her x-1 33-37 [1] 1.700 [2] 1.56 3.1 smc x-1 50-70 [3] 3.89 [4] 11.0 8.9 lmc x-4 30 [5] 1.41 [6] 10.6 4.5 cyg x-2 60-90 [7] 9.844 [8] 0.34 8.0 x1820-303 171 [9] 0.008 [10] [0.1] 0.1 ss433 162 [11] 13.10 [12] [1.0] [3.0] x1916-053 5 [13] 0.035 [14],[15] [0.1] 0.2 199 [16] [1] leahy & igna (2010), [2] tananbaum et al. (1972), [3] clarkson et al. (2003), [4] schreier et al. (1972), [5] lang et al. (1981), [6] chevalier & ilovaisky 1977, [7] clarkson et al. (2003), [8] cowley et al. (1979), [9] chou & grindlay (2001), [10] stella et al. (1987), [11] margon (1984), [12] crampton et al. (1980), [13] homer et al. (2001), [14] walter et al. (1982), [15] white & swank (1982), [16] priedhorsky & terrell (1984) figure 2: her x-1: this steady super-orbital period have been associated with a stable radiation-driven warp. 129 m. m. kotze, p. a. charles figure 3: smc x-1: the super-orbital period that is evolving dramatically over an even longer time-scale, may be associated with a radiation-driven warp that is variable. figure 4: lmc x-4: the steady long-term period may be produced by a stable radiation-driven warp. figure 5: cyg x-2: the unstable/chaotic behaviour may be associated with chaotic warping. 3.2 third body zdziarski et al. (2007) performed a comprehensive analysis of the triple system scenario and how binary eccentricity oscillations could modulate mass transfer through l1, finding agreement between theoretical and observed lightcurves for x1820-303. figure 6: x1820-303. 3.3 precessing relativistic jets ss433 is the prototypical microquasar. its precessing relativistic jets and their resulting ∼ 162 day superorbital period was obtained by hjellming et al. (1981) and margon (1984) from variations in the radial velocity measurements from balmer and he i emission lines. figure 7: ss433. 3.4 tidally-induced disc precession whitehurst & king (1991) described how tidal interactions with the donor may excite resonances in the accretion disc, causing it to precess and produce quasi130 investigating long-term behaviour of x-ray binaries using archival data periodic variations in the lightcurve. essentially the disc becomes elliptical, expanding beyond its critical radius while remaining within its roche-lobe radius, so that disc precession is effectively the result of changes in the orientation of this elliptical accretion disc with respect to the donor. figure 8: x1916-053. tidal disc precession depends on the mass ratio (q = m2 m1 ) and will only occur if q < 0.25 − 0.33, producing so-called “superhumps”, which have been detected in the su uma sub-class of cataclysmic variables (cvs). the vertical dashed line on fig. 1 indicates this boundary, to the left of which xrbs are expected to be susceptible to tidal disc precession (e.g. x1916-053 and x1820-303). 4 discussion the composite figure for each source contains a wealth of information to assist interpretation. the lightcurves are shown in the top panels and density maps with the dps results form the main panels, for which the scales are indicated to the right. the 99.9%-confidence white noise levels (green lines), together with the l-s for the entire datasets, are shown in the left panels. all included sources show significant (exceeding the white noise levels) periodic signals in their dps. previously reported periodicities (listed in table 1), indicated by red bars, are clearly present in the dps panels and often also in the l-s (left-hand) panels along with additional variability. 5 conclusions quasi-periodic or aperiodic signals should be investigated using time-dependent period analysis techniques. while these examples focussed on long-term superorbital behaviour of xbs, it should be immediately apparent how useful such a technique would be for investigating short-term quasi-periodic oscillations in cvs. acknowledgement asm results were provided by the asm/rxte teams at mit and at the rxte science operations facility and guest observer facility at nasa’s goddard space flight center (gsfc). this work formed part of my phd thesis, supervised by prof. phil charles and co-supervised by prof. brian warner, at the university of cape town (uct) and was funded by the national research foundation (nrf) of south africa. references [1] chevalier c., ilovaisky s. a. : 1977, a&a, 59, l9 [2] chou y., grindlay j. e. : 2001, apj, 563, 934 doi:10.1086/324038 [3] clarkson, w. i., charles, p. a., coe, m. j., laycock, s. : 2003, mnras, 343, 1213 doi:10.1046/j.1365-8711.2003.06761.x [4] cowley a. p., crampton d., hutchings j. b. : 1979, apj, 231, 539 doi:10.1086/157216 [5] crampton d., cowley a. p., hutchings j. b. : 1980, apj, 235, l131 doi:10.1086/183176 [6] hjellming, r. m., & johnston, k. j. : 1981, nature, 290, 100 doi:10.1038/290100a0 [7] homer l., charles p. a., hakala p., muhli p., shih i.-c., smale a. p., ramsay g. : 2001, mnras, 322, 827 [8] kotze, m. m., charles, p. a. : 2012, mnras, 420, 1575 doi:10.1111/j.1365-2966.2011.20146.x [9] lang f. l., et al. : 1981, apj, 246, l21 [10] leahy d. a., igna c. d. : 2010, apj, 713, 318 doi:10.1088/0004-637x/713/1/318 [11] levine, a. m., bradt, h., cui, w., et al. : 1996, apjl, 469, l33 doi:10.1086/310260 [12] lomb, n. r. : 1976, apss, 39, 447 [13] margon, b., anderson, s. f., aller, l. h., downes, r. a., & keyes, c. d. : 1984, apj, 281, 313 [14] ogilvie, g. i., & dubus, g. :2001, mnras, 320, 485 131 http://dx.doi.org/10.1086/324038 http://dx.doi.org/10.1046/j.1365-8711.2003.06761.x http://dx.doi.org/10.1086/157216 http://dx.doi.org/10.1086/183176 http://dx.doi.org/10.1038/290100a0 http://dx.doi.org/10.1111/j.1365-2966.2011.20146.x http://dx.doi.org/10.1088/0004-637x/713/1/318 http://dx.doi.org/10.1086/310260 m. m. kotze, p. a. charles [15] priedhorsky w. c., terrell j. : 1984, apj, 280, 661 [16] scargle, j. d. : 1982, apj, 263, 835 [17] schreier e., giacconi r., gursky h., kellogg e., tananbaum h. : 1972, apj, 178, l71 doi:10.1086/181086 [18] stella l., priedhorsky w., white n. e. : 1987, apj, 312, l17 doi:10.1086/184811 [19] tananbaum h., gursky h., kellogg e. m., levinson r., schreier e., giacconi r. : 1972, apj, 174, l143 [20] walter f. m., mason k. o., clarke j. t., halpern j., grindlay j. e., bowyer s., henry j. p. : 1982, apj, 253, l67 [21] whitehurst, r., & king, a. : 1991, mnras, 249, 25 [22] white n. e., swank j. h. : 1982, apj, 253, l61 doi:10.1086/183737 [23] zdziarski, a. a., pooley, g. g., & skinner, g. k.: 2011, mnras, 412, 1985 discussion koji mukai: can you clarify what you saw in x1916? it is a low q system, but i thought the disc precessesion period was about 5 days, while your plot showed only longer periods. marissa kotze: our dps analysis for this source show no significant periodic behaviour below 10 days. but this is not unexpected, since the 5 day period associated with tidal disc precession in x1916-053 was not determined from x-ray flux variations, but rather from optical variability and changes in the structure of the x-ray dips, which our dps analysis of rxte asm data is unable to detect. 132 http://dx.doi.org/10.1086/181086 http://dx.doi.org/10.1086/184811 http://dx.doi.org/10.1086/183737 introduction data analysis rxte/asm archival data time-dependent period analysis results radiation-induced disc warp/tilting third body precessing relativistic jets tidally-induced disc precession discussion conclusions 50 acta polytechnica ctu proceedings 2(1): 50–54, 2015 50 doi: 10.14311/app.2015.02.0050 optical photometry of lmxbs: uw crb (=ms 1603+260) and v1408 aql (=4u 1957+115) p. a. mason1,2, e. l. robinson3, s. gomez1, j. v. segura1 1department of physics, university of texas at el paso, 500 w. university, el paso, tx, usa, 79902 2department of mathematics and physical sciences, new mexico state university dacc, las cruces, nm, usa, 88003 3department of astronomy, university of texas at austin, 1 university station c1400, austin, tx, usa, 78712 corresponding author: pmason@nmsu.edu abstract we present new optical observations of v1408 aql (= 4u 1957+115), the only low mass x-ray binary, black hole candidate known to be in a persistently soft state. we combine new broadband optical photometry with previously published data and derive a precise orbital ephemeris. the optical light curves display sinusoidal variations modulated on the orbital period as well as large night to night changes in mean intensity. the amplitude of the variations increases with mean intensity while maintaining sinusoidal shape. considering the set of constraints placed by the x-ray and optical data we argue that v1408 aql may harbor a very low mass black hole. optical light curves of uw crb display partial eclipses of the accretion disk by the donor star that vary both in depth and orbital phase. the new eclipses of uw crb in conjunction with published eclipse timings are well fitted with a linear ephemeris. we derive an upper limit to the rate of change of the orbital period. by including the newly observed type i bursts with published bursts in our analysis, we find that optical bursts are not observed between orbital phases 0.93 and 0.07, i.e. they are not observable during partial eclipses of the disk. keywords: low mass x-ray binaries photometry individual: v1408 aql individual: uw crb. 1 introduction low mass x-ray binaries (lmxbs) are composed of a low mass donor and a compact object, either a black hole or a neutron star primary. the compact star accretes matter from the donor via an accretion disk. we present updates on our previous work on the black hole candidate v1408 aql and the lmxb containing a neutron star, uw crb (see mason et al. 2012). 2 v1408 aql v1408 aql is a lmxb whose secondary star fills its roche lobe and the compact object is a black hole candidate. this is the only lmxb black-hole candidate known to be in a persistently active soft state. there are only three other x-ray binaries that have been found to be persistently active, lmc x-1, lmc x-3 and cyg x1. but unlike v1408 aql, these are all high mass x-ray binaries. v1408 aql has always remained in the spectrally soft, disk dominated, x-ray state, never reaching the low/hard state. the persistent activity of v1408 aql and the disk-dominated soft x-ray spectrum are likely to be related. 2150 2151 2152 2153 2154 2155 0.02 0.025 0.03 0.035 0.04 0.045 0.05 hjd − 2454000 r e la ti v e i n te n s it y figure 1: optical photometry of v1408 aql on four nights during 10-15 august 2012. intensity is shown relative to the combined light of stars labeled 6 and 8 of figure 1 in bayless et al. (2011). an orbital sine wave like variation is seen each night as well as a large drop in brightness between the second and third night. once night to night variations are removed, the light curve is nearly sinusoidal. the optical light curve of v1408 aql shows a sinusoidal orbital modulation at the system’s 9.33 hr orbital 50 http://dx.doi.org/10.14311/app.2015.02.0050 optical photometry of lmxbs: uw crb (=ms 1603+260) and v1408 aql (=4u 1957+115) period. this period was first measured by thorstensen (1987) and it was later confirmed by bayless et al. (2011) and mason et al. (2012). we obtained new high-speed optical photometry of v1408 aql using the 2.1m otto struve telescope of mcdonald observatory on four nights in august 2012 and five nights in july 2012. photometry was obtained using the argos ccd photometer (mukadam and nather, 2005) and a bvr (broad-band) filter with a time resolution of 10 seconds. the light curves from august 2012 are shown in figure 1. strong night to night transitions such as these are often observed. as part of our analysis, nightly mean variations were removed and normalized to form a data set for period analysis. once normalized and phased the orbital modulation becomes apparent, and a clear sinusoidal modulation is observed. 2.1 updated ephemeris for v1408 aql using all the available optical photometry for a total of 126 hours of observation spread over 29 nights and 5 years, we construct a phase dispersion minimization periodogram (stellingwerf, 1978), see figure 2. preliminary results were presented by gomez et al. (2013). we derive a precise orbital period of 0.388893(3) days. this improved period is consistent with previously published results. figure 2 also shows previous period determinations with error bars (thorstensen 1987; bayless et al. 2011; mason et al. 2012). the improved ephemeris for the time of maximum flux is t (hjd) = 2454621.829(4) + 0.388893(3)e 0.15 0.2 0.25 0.3 0.35 0.4 0.38820.38840.38860.38880.3890.38920.3894 p h a s e d is p e rs io n period (days) thorstensen, 1987 current period mason, 2012 bayless, 2011 figure 2: period search results using phase dispersion minimization. previously published results are also shown. 2.2 the nature of the compact star multi-wavelength datasets yield a variety of constraints on the compact component of this binary. the x-ray data show a soft spectrum well described by a multitemperature blackbody and in 15% of the observations an additional non-thermal steep power law component is seen (nowak et al. 2012). the optical intensity also occasionally increases sharply; see the histogram of nightly mean brightness in figure 3. the distance of v1408 aql is likely d > 7 kpc, but the compact star mass remains largely unconstrained by x-ray data (nowak et al. 2012). radio emission is 330-810 times fainter than expected in relation to other x-ray binaries (russell et al. 2011). 0 0.005 0.01 0.015 0.02 0.025 0 2 4 6 8 10 12 14 16 increase from minimum intensity c o u n ts ( n ig h ts ) figure 3: histogram of nightly optical intensity variations. in order to examine the sinusoidal modulation, light curves were shifted by the amounts shown in the histogram. zero on this linear scale is the relative intensity of the minimum observed nightly mean. for example, the first night shown in figure 1 is shifted down by approximately 0.013 to remove the night to night variation. on most nights v1408 aql remained near minimum intensity, while on one night (14 july 2012) it showed a particularly high intensity. the optical data suggest that the orbital modulation is dominated by the heated face of the donor star. even as strong night to night intensity fluctuations occur, a strong sinusoidal modulation remains at all brightness levels. bayless et al. (2011) estimated a mass ratio of q = m2/m1 > 0.3. the primary is most likely a black hole, since the system has never shown a type-i x-ray burst since its discovery in 1973, and the black hole candidate has a fast spin and a spectrum typical for black holes (nowak et al. 2008). fits of models to the optical light curve give strong evidence for a low mass primary with a mass of m < 7m�, as well as a low inclination of i = 11o − 25o. in summary, a low mass black hole model is consistent with both optical and x-ray emission studies of v1408 aql. 3 uw crb the lmxb uw crb contains a neutron star primary. the optical light curve displays partial eclipses of the disk by the donor and frequent bursts (mason et al. 51 p. a. mason et al. 2008, 2012, hakala et al. 2009). the appearance of type i x-ray bursts, either as x-rays or reprocessed as optical emission clearly indicates that the compact object is a neutron star. the secondary star is likely significantly evolved. an hst/cos uv spectrum uw crb was obtained (froning et al. 2012) in which the c iv lines are totally absent. we obtained optical photometry at the mcdonald observatory otto struve 2.1-m telescope in june 2013, using the argos ccd photometer (mukadam and nather, 2005), a bvr filter, and 10-second time resolution (see figure 4). several bursts are easily noticed. the burst frequency varies from night to night. 0.00 0.05 0.10 0.15 0.20 0.0 1.0 2.0 3.0 r e la � v e i n te n s it y orbital phase 2013, june 10 2013, june 11 2013, june 12 2013, june 13 figure 4: optical photometry of uw crb on four successive nights in june 2013 is shown, compared to star c in the finder chart of hakala et al. (1998). the light curves taken on the first three nights have been shifted in intensity relative to the fourth by increments of 0.05. the phase and shape of the partial eclipses fluctuate over a 5.5 day beat cycle. see mason et al. (2008, 2012) and hakala et al. (2009). 3.1 an updated ephemeris for uw crb we measured 7 new eclipse times which we combined with 56 eclipses listed by mason et al. (2008, 2012) for a total of 63 eclipses. the new best-fit linear ephemeris is: t (hjd) = 24553118.83822(47) + 0.077067211(17)e where e is the eclipse number. we also fit a quadratic function to the eclipses in the o-c diagram (see figure 5). the quadratic term is not significant, but allows the derivation of an upper limit to the period derivative of (at 90 % confidence): |dp/dt| < 3.0 × 10−11s/s. 3.2 optical bursts as a mixture of hydrogen and helium accumulates on the neutron star, some hydrogen fuses steadily while a he-rich layer builds up. when this atmosphere becomes dense and hot enough, it suddenly fuses into carbon and emits a type i x-ray burst (e.g. strohmayer and bildsten, 2006). 0 -0.02 -0.01 0 0.01 0.02 figure 5: o-c diagram for 63 eclipses of uw crb with respect to the best-fit linear ephemeris. the dashed line is the best-fit quadratic function. however, the quadratic term is not statistically significant. the vertical scatter of the points is due to periodic phase shifting of the eclipses. figure 6: photometry of a burst that occurred on june 12, 2013. the intensity, compared to star c in figure 1 of hakala et al. (1998), of this typical burst rises to peak in 10-20 s and declines in about 50 s. uw crb is a frequent x-ray burster (hakala et al., 2005). optical bursts are observed (e.g. hynes, robinson, and jeffery, 2004) because regions in the accretion disk absorb x-rays and are heated. the heated areas emit optical radiation and we see a burst while the optical emission regions cool. an example optical burst is shown in figure 6. when combined with published timings (see mason et al. 2012 and hakala et al. 2009 and references therein) 40 bursts events have been ob52 optical photometry of lmxbs: uw crb (=ms 1603+260) and v1408 aql (=4u 1957+115) served. these are shown as a function of orbital phase in figure 7. notice that there is a partial eclipse zone where no bursts have been detected. if the bursts were being reprocessed off the secondary star, the number and strength of the bursts would have decreased slowly before and increased slowly after eclipses. instead, the number drops abruptly at the edge of the eclipse. this is strong evidence that the optical burst emission region is the disk and not the heated face of the secondary. figure 7: distribution of 40 type i bursts as a function of phase. no bursts occurred between phase 0.93 and 1.07 due to partial eclipses of the accretion disk by the companion. 4 conclusions a new ephemeris is derived for v1408 aql. the light curves show sinusoidal modulation with night-to-night changes in mean brightness. our models favor a black hole primary with a mass less than 7 solar masses. we suggest that the orbital modulation is caused primarily by the changing aspect of the heated face of the donor star. v1408 aql (= 4u 1957+115) is likely a black hole, and there is increasing evidence that it is near the lower end of the observed black hole mass distribution. analysis of new and previously published partial eclipses of the accretion disk in uw crb are well fitted with a linear ephemeris. in addition, there is ∼ 5 day periodic shifting around phase 0 due to precession of an elliptical accretion disk (mason et al. 2008, 2012; hakala et al. 2012). an upper limit on the rate of change of the orbital period based on the best fit quadratic ephemeris is derived, see equation 1. strong bursts are seen at all phases outside of eclipse indicating that the disk is the source of the re-processing responsible for the optical bursts, rather than the heated face of the donor star. acknowledgement we thank tom maccarone for discussions concerning the nature of the compact object in v1408 aql (=4u 1957+115). we thank the referee, hans ritter, for a careful review and many helpful comments and corrections. this work is supported by nsf/paare grant no. 0958783. references [1] bayless, a. j., robinson, e. l., mason, p. a., robertson, p., 2011, apj, 730, 43. doi:10.1088/0004-637x/730/1/43 [2] froning, c. s., 2012, aas, 219, 153.12. [3] gomez, s., mason, p. a., robinson, e. l., 2013, aas, 221, 142.25. [4] hakala, p., et al., 2005, mnras, 356, 1133. [5] hakala, p., hjalmarsdotter, l., hannikainen, d. c., muhli, p., 2009, mnras, 394, 892. [6] hynes, r. i., robinson, e. l., jeffery, e., 2004, apj, 608, 101. doi:10.1086/422471 [7] mason, p. a., robinson, e. l., gray, c. l., hynes, r. i., 2008, apj, 685, 428. doi:10.1086/590381 [8] mason, p. a., robinson, e. l., bayless, a. j., hakala, p. j., 2012, aj, 144, 108. [9] mukadam, a. s., nather, r. e., 2005, japa, 26, 321. [10] nowak, m. a., et al., 2008, apj, 689, 1199. doi:10.1086/592227 [11] nowak, m. a., wilms, j., pottschmidt, k., schulz, n., maitra, d., miller, j., 2012, apj, 744, 107. doi:10.1088/0004-637x/744/2/107 [12] russell, d. m., et al., 2011, apj, 739, l19. doi:10.1088/2041-8205/739/1/l19 [13] stellingwerf, r. f., 1978, apj, 224, 953. [14] strohmayer, t., bildsten, l., 2006, in ”compact stellar x-ray sources”, ed. by w. lewin and m. van der klis, (cambridge: cambridge univ. press) p. 113. doi:10.1017/cbo9780511536281.004 [15] thorstensen, j. r., 1987, apj, 312, 739. discussion simone scaringi: could the absence of type i x-ray bursts at specific phases be used to constrain the size of the secondary star and/or the emitting region giving rise to the burst. 53 http://dx.doi.org/10.1088/0004-637x/730/1/43 http://dx.doi.org/10.1086/422471 http://dx.doi.org/10.1086/590381 http://dx.doi.org/10.1086/592227 http://dx.doi.org/10.1088/0004-637x/744/2/107 http://dx.doi.org/10.1088/2041-8205/739/1/l19 http://dx.doi.org/10.1017/cbo9780511536281.004 p. a. mason et al. paul mason: yes, especially when combined with information about the mass ratio. currently, we do constrain the optical burst emission region to the disk. we exclude significant re-processing off of the companion star. 54 introduction v1408 aql updated ephemeris for v1408 aql the nature of the compact star uw crb an updated ephemeris for uw crb optical bursts conclusions 108 acta polytechnica ctu proceedings 1(1): 108–112, 2014 108 doi: 10.14311/app.2014.01.0108 mass accretion processes in young stellar objects: role of intense flaring activity salvatore orlando1, fabio reale2,1, giovanni peres2,1, andrea mignone3 1inaf osservatorio astronomico di palermo, piazza del parlamento 1, 90134, palermo, italy 2dip. di fisica e chimica, università degli studi di palermo, piazza del parlamento, 1, 90134, palermo, italy 3dip. di fisica generale, università degli studi di torino, via pietro giuria 1, 10125, torino, italy corresponding author: orlando@astropa.inaf.it abstract according to the magnetospheric accretion scenario, young low-mass stars are surrounded by circumstellar disks which they interact with through accretion of mass. the accretion builds up the star to its final mass and is also believed to power the mass outflows, which may in turn have a significant role in removing the excess angular momentum from the star-disk system. although the process of mass accretion is a critical aspect of star formation, some of its mechanisms are still to be fully understood. on the other hand, strong flaring activity is a common feature of young stellar objects (ysos). in the sun, such events give rise to perturbations of the interplanetary medium. similar but more energetic phenomena occur in ysos and may influence the circumstellar environment. in fact, a recent study has shown that an intense flaring activity close to the disk may strongly perturb the stability of circumstellar disks, thus inducing mass accretion episodes (orlando et al. 2011). here we review the main results obtained in the field and the future perspectives. keywords: accretion, accretion disks mhd stars: circumstellar matter stars: flare stars: pre-main-sequence x-rays: stars. 1 introduction observations in the x-ray band reveal that low-mass pre-main-sequence stars are strong sources with x-ray luminosities 3−4 orders of magnitude greater than that of the present-day sun. the source of this high-energy radiation is plasma with temperatures of 1 − 100 mk in the stellar outer atmospheres (coronae), heated by magnetic activity analogous to the solar one but higher by factors up to 106. such a magnetic activity manifests through very different phenomena that may occur in several places of the stellar atmosphere and circumstellar environment. the young star interacts with its disk in a complex fashion, with accretion and ejection of collimated outflows. strong magnetic fields are believed to connect the star with a keplerian circumstellar disk, funneling accretion onto limited portions of the stellar surface (e.g. hartmann 1998) where shocks are produced by the impact of the accretion streams (e.g. orlando et al. 2010). x-ray flares are violent manifestations of the stellar magnetic activity and are triggered by an impulsive energy input from coronal magnetic fields. x-ray observations in the last decades have shown that flares in young stellar objects (ysos) have amplitudes much larger than solar analogues and occur much more frequently. examples of these flares are those collected by the chandra satellite in the orion star-formation region (coup enterprise; favata et al. 2005). in the sun, such energetic events are often associated to coronal mass ejections and give rise to perturbation effects of the interplanetary medium, broadly known as space weather effects. similar phenomena are expected to occur in young stars, and may affect the circumstellar environment. for instance, favata et al. (2005) analyzed the most energetic flares observed by coup and found that these flares might be hosted in magnetic loops extending several stellar radii, much larger than ever observed in older stars. since the central star is surrounded by a circumstellar disk accreting material onto the star, it is natural to ask whether strong flaring activity involves the disk and even perturbs its stability, possibly affecting the mass accretion to the star. at the present time, in fact, it is unclear where these flares occur. the differential rotation of the disk together with the interaction of the disk with the magnetosphere may cause magnetic reconnection close to the disk’s surface, triggering large-scale flares there. in this case, the flares may perturb the stability of the circumstellar disk causing, in particular, a strong local overpressure. the pressure gradient force might be able to 108 http://dx.doi.org/10.14311/app.2014.01.0108 mass accretion processes in young stellar objects: role of intense flaring activity push disk’s matter out of the equatorial plane into funnel streams, thus providing a mechanism to drive mass accretion that differs from that, commonly invoked in the literature, based on the disk viscosity (romanova et al. 2002). bright flares close to circumstellar disks may therefore have important implications for a number of issues such as the transfer of angular momentum and mass between the star and the disk, the powering of outflows, and the ionization of circumstellar disks. recently, we have investigated the idea that an intense flare close to an accretion disk may perturb the stability of the disk and trigger mass accretion onto the star (orlando et al. 2011). in this paper we review our findings and present preliminary results of a study investigating the effects of a storm of small-to-medium flares on the stability of accretion disks (orlando et al. 2014). in sect. 2 we describe the mhd model; in sect. 3 we present the results; in sect. 4 we draw our conclusions. 2 the mhd model the model describes a classical t tauri star (ctts) of mass m∗ = 0.8m� and radius r∗ = 2r� located at the origin of a 3d spherical coordinate system (see orlando et al. 2011 for a detailed description). we adopted the initial conditions introduced by romanova et al. (2002). in particular, we assumed the rotation period of the star to be 9.2 days as representative of cttss. the initial unperturbed stellar atmosphere is approximately in equilibrium and consists of three components: the stellar magnetosphere, the extended stellar corona, and the keplerian disk. initially, the magnetosphere is assumed to be force-free, with dipole topology and magnetic moment aligned with the rotation axis of the star. the magnetic moment is chosen in order to have a magnetic field strength of the order of 1 kg at the stellar surface. the isothermal disk is cold, dense and rotates with angular velocity close to the keplerian value; its rotation axis (coincident with the rotation axis of the star) is aligned with the magnetic moment. the disk is initially truncated by the stellar magnetosphere at the radius rd where the ram pressure of the disk is equal to the magnetic pressure; for the adopted parameters, rd = 2.86r∗ and the co-rotation radius is located at rco = 9.2r∗. the corona is initially isothermal with temperature t = 4 mk and at low density. the system is described by the time-dependent mhd equations of mass, momentum, and energy conservation in a 3d spherical coordinate system, extended with gravitational force, viscosity of the keplerian disk, thermal conduction (including the effects of heat flux saturation), coronal heating (via a phenomenological term), and radiative losses from optically thin plasma (see orlando et al. 2011 for more details). the phenomenological heating is prescribed as a component, describing the stationary coronal heating, plus a transient component, triggering the flares. the calculations were performed using pluto (mignone et al. 2007), a modular, godunov-type code for astrophysical plasmas. 3 results as a first application of the model, we have investigated the effects of a single bright flare on the stability of the disk (orlando et al. 2011). the initial heat pulse triggering the flare is released at a distance of 5r∗, namely in a region comprised between the truncation and corotation radii, and it is supposed to be indicative of a likely area of magnetic reconnection. heat pulses occurring closer to the inner disk edge are expected to produce analogous perturbations on the disk dynamics. we followed the evolution of the star-disk system for ≈ 2 days. after the heat pulse has been released, an mhd shock wave develops above the disk and propagates away from the protostar. the heat deposition determines a local increase of temperature and pressure. disk material is heated and expands with a strong evaporation front which is channeled along the magnetic field lines towards the central star. after ∼ 1 hour a hot loop forms linking the disk with the star (see upper panel in fig. 1). the loop has an effective maximum temperature of ≈ 100 mk (at the peak of emission measure) and a length of ≈ 1012 cm; these values are in good agreement with those inferred from the analysis of the brightest flares observed by coup (favata et al. 2005). the overheating of the disk surface makes a significant amount of material expand and be ejected in the magnetosphere. a small fraction of this material fills the loop whose density increases from 108 cm−3 to 1010 cm−3. on the other hand, most of the expanding disk material is not confined by the magnetic field and is ejected away from the star, carrying away mass and angular momentum. during the evolution of the hot loop, an overpressure wave develops where the heat pulse has been injected. this overpressure travels through the disk and reaches the opposite boundary after ≈ 5 hours. there the pressure gradient force drives the material out of the disk and channels it into a funnel flow. then the gravitational force accelerates the escaped material toward the central star where the stream impacts ≈ 25 hours after the injection of the heat pulse. the accretion flow persists until the end of the simulation (t = 48 h). the lower panel in fig. 1 shows a cutaway view of the stardisk system after the impact of the stream onto the stellar surface. we found that the dynamics and physical characteristics of the accretion stream triggered by the flare closely recall those of streams driven by the accumulation of mass at the disk truncation radius under 109 salvatore orlando et al. the effect of the viscosity and pushed out of the equatorial plane because of the growing pressure gradient there (e.g. romanova et al. 2002). figure 1: effects of a single bright flare on the stability of the circumstellar disk. cutaway views of the star-disk system showing the mass density of the disk (yellow-green) at t = 2.0 hours (upper panel) and at t = 45 hours (lower panel) since the injection of the heat pulse. the upper panel also over-plots the threedimensional volume rendering of the plasma temperature (mk), showing the flaring loop (in red) linking the inner part of the disk with the star. the lower panel shows the accretion stream triggered by the flare in the side of the disk opposite to the flaring loop. selected magnetic field lines are overplotted in red. from the simulation, we derived also the mass accretion rate and found ṁ > 2.5 × 10−10 m� yr−1. we compared this rate with those inferred from optical observations of cttss (herczeg & hillenbrand 2008; curran et al. 2011) and found a good agreement. we concluded that a bright flare as those frequently detected in ysos (e.g. coup observations) can be an efficient mechanism to trigger accretion onto the protostar itself with accretion rates on the same order of those commonly measured in cttss. as a follow-up of the previous study, we explored in more details the possibility that significant mass accretion in young stars can be triggered by a storm of small-to-medium flares (as those frequently observed) occurring on the accretion disk (orlando et al. 2014). to this end, we performed a 3d mhd simulation analogous to that described in orlando et al. (2011) but considering a storm of flares distributed randomly in proximity of the disk surface instead of a single bright flare. we found that each simulated flare follows an evolution similar to that of the single bright flare described in orlando et al. (2011). the main difference is that, now, interactions between next flares may occur. figure 2 shows cutaway views of the star-disk system after ≈ 26 hours. during the system evolution, the flares build up an extended corona linking the star with the disk. at the same time, the disk is strongly perturbed by the flares and, after ≈ 20 hours, several funnel streams develop, accreting substantial mass onto the star (see lower panel in fig. 2). the streams persist until the end of the simulation and last for a time longer than the typical interval between flares. the simulated mass accretion rate is 10−10 < ṁ < 10−9 m�yr −1, again in good agreement with the values inferred from optical observations of cttss (herczeg & hillenbrand 2008; curran et al. 2011). 4 conclusions we investigated the effects of an intense flaring activity on the stability of a circumstellar disk surrounding a magnetized ctts. to this end, we developed a 3d mhd model including, for the first time, all key physical processes, most notably the thermal conduction and the radiative losses from optically thin plasma. as a first step, we analyzed the perturbation induced by a single bright flare occurring in proximity of the disk (orlando et al. 2011). then, we investigated the effects of a storm of small-to-medium flares distributed randomly close to the disk (orlando et al. 2014). our findings lead to the following conclusions: (a) flares occurring close to the circumstellar disk can trigger substantial and persistent accretion flows, similar to those caused by the disk viscosity; (b) an intense flaring activity close to the disk builds up an extended corona linking the star to the disk. in the case of a single bright flare occurring close to the disk, the simulations suggest that mass accretion events associated with x-ray flares should be observed. however, correlation between uv/optical accretion tracers and x-ray flux is rarely seen (stassun et 110 mass accretion processes in young stellar objects: role of intense flaring activity al. 2006). on the other hand, such a correlation is not prediced by simulations describing a continuous flaring activity close to the disk (orlando et al. 2014). in this case, the streams are triggered by the first flares at the beginning of the simulation and then are continuously powered by the following ones. at regime, therefore, no clear correlation between uv/optical accretion tracers and x-ray flux is foreseen. in the light of the above findings, we suggest that the flaring activity common to ysos may turn out to be important in the exchange of angular momentum and mass between the circumstellar disk and the central protostar. figure 2: effects of a storm of flares on the disk stability. cutaway views of the star-disk system showing the mass density of the disk (blue) at t = 26 hours. the panels over-plot the three-dimensional volume rendering of the plasma temperature (in mk), showing the flaring loops (in red) linking the inner part of the disk with the star. the upper panel shows the star-disk system observed edge-on, whereas the lower panel shows the system observed pole-on. selected magnetic field lines are overplotted in the upper panel. acknowledgement we thank the referee for constructive and helpful criticism. pluto is developed at the turin astronomical observatory in collaboration with the department of general physics of the turin university. we acknowledge the cineca award n. hp10a4zcv5,2012 for the availability of high performance computing resources and support. references [1] curran, r.l., argiroffi, c., sacco, g.g., orlando, s., peres, g., reale, f., maggio, a.: 2011, a&a 526, a104 [2] favata, f., flaccomio, e., reale, f., micela, g., sciortino, s., shang, h., stassun, k.g., feigelson, e.d.: 2005, apjs 160, 469 doi:10.1086/432542 [3] hartmann l.: 1998, accretion processes in star formation. new york: cambridge university press; cambridge astrophysics series; 32 [4] herczeg, g.j., hillenbrand, l.a.: 2008, apj 681, 594 doi:10.1086/586728 [5] mignone, a., bodo, g., massaglia, s., matsakos, t., tesileanu, o., zanni, c., ferrari, a.: 2007, apjs 170, 228 doi:10.1086/513316 [6] orlando, s., sacco, g.g., argiroffi, c., reale, f., peres, g., maggio, a.: 2010, a&a 510, a71 [7] orlando, s., reale, f., peres, g., mignone, a.: 2011, mnras 415, 3380. [8] orlando, s., reale, f., peres, g., mignone, a.: 2014, in preparation [9] romanova, m.m., ustyugova, g.v., koldoba, a.v., lovelace, r.v.e.: 2002, apj 578, 420 doi:10.1086/342464 [10] stassun, k.g., van den berg, m., feigelson, e., flaccomio, e.: 2006, apj 649, 914 doi:10.1086/506422 discussion matteo guainazzi: what kind of x-ray variability does your model predict? is it observed in the coup sources? salvatore orlando: the x-ray variability depends on the rate of flares triggered. in the simulation with a single flare, we observe the fast rise and subsequent slow decay of x-ray emission characteristic of flares. in the simulation with a storm of flares, we observe a background emission due to the many small flares evolving simultaneousluy with superimposed local peaks due to the brightest flares. the x-ray variability found is consistent with that of ysos observed by coup. 111 http://dx.doi.org/10.1086/432542 http://dx.doi.org/10.1086/586728 http://dx.doi.org/10.1086/513316 http://dx.doi.org/10.1086/342464 http://dx.doi.org/10.1086/506422 salvatore orlando et al. david buckley: do you believe that your magnetospheric disk interaction models could be extrapolated to intermediate polar class of cvs, where accretion occurs from disrupted disc onto a magnetic white dwarf? salvatore orlando: our model can be extrapolated (with appropriate scaling) to systems in which mass accretion occurs from a circumstellar disk, if relativistic effects can be neglected. 112 introduction the mhd model results conclusions 303 acta polytechnica ctu proceedings 2(1): 303–307, 2015 303 doi: 10.14311/app.2015.02.0303 gloria and study of cataclysmic variables r. hudec1,2, v. šimon1,2 1astronomical institute, academy of sciences of the czech republic, cz-25165 ondřejov, czech republic 2czech technical university in prague, faculty of electrical engineering, prague, czech republic corresponding author: rene.hudec@asu.cas.cz abstract we report here on the ongoing eu fp7 project gloria (global robotic-telescopes intelligent array) with emphasis on possibility of investigation of cataclysmic variables by users. gloria will enable the first free and open-access network of robotic telescopes in the world. we show several examples of the not often used topics (but suitable for gloria) for the studies of activity of cataclysmic variables, e.g. search for outbursts in intermediate polars and the fluctuations of brightness in their quiescence, and investigation of the optical counterparts of supersoft x-ray sources. keywords: catalysmic variables: optical photometry spectroscopy: low-dispersion spectra data mining robotic telescopes project eu fp7 gloria. 1 introduction the research in astronomy poses two main challenges, namely 1) the immensity of the sky and 2) the huge amount of astronomical data being gathered. in fact, astronomers are nowadays facing great difficulty in finding the resources to analyze the increasing flood that modern astronomy instruments generate,insufficient computing power, insufficiently powerful software tools, and not enough hours in the day. the sky comprises 40 000 square degrees (144 million square minutes). the future professional projects (like lsst) intend to observe a very significant fraction of it on a regular basis. in order to meet the abovementioned challenges, an increasing number of astronomical projects have begun to try to foster citizen participation to help analyze data by using collaborative internet application (the so-called web 2.0). gloria stands for “global robotic-telescopes intelligent array”. gloria will be the first free and open-access network of robotic telescopes in the world. it represents a web 2.0 environment where users can do research in astronomy by observing with robotic telescopes, and/or by analyzing data that other users have acquired with gloria, or from other free access databases like the european virtual observatory (http://www.euro-vo.org). the gloria project (http://gloria-project.eu) represents the first attempt to create a network of about 20 robotic telescopes for public access, education, science and much much more. the gloria network has 17 network telescopes now. few others are expected to join us soon. many internet communities have already formed to speed-up scientific research, to collaborate in documenting something, or for social projects. research in astronomy can benefit from attracting many eyes to the sky – to detect something in the sky requires looking at the right place at the right time. our robotic telescopes can search the sky, but the vast quantities of data they produce are far greater than astronomers have time to analyze. gloria will provide a way of putting thousands eyes and minds on an astronomy problem. gloria is intended to be a web 2.0 structure, with the possibility of doing real experiments. the community will not only generate content, as in most web 2.0, but it will control telescopes around the world, both directly and via scheduled observations. the community will take decisions for the network and it will give “intelligence” to gloria, while the drudge work (such as drawing up telescope schedules that satisfy various constraints) will be done by algorithms that will be developed for the purpose. gloria project will define free standards, protocols and methodology for the following purposes: controlling robotic telescopes and all related instrumentation (i.e. cameras, filter-wheels, domes...); giving web access to the network: access to an arbitrary number of robotic telescopes via a web portal; conducting online experiments (it will be possible to design specific web environments for controlling telescopes for research on a specific scientific issue; conducting off-line experiments (it will be possible to design specific web environments for analyzing astronomical meta-data produced by gloria or other databases). 303 http://dx.doi.org/10.14311/app.2015.02.0303 r. hudec, v. šimon the main part of gloria is represented by providing access to robotic telescopes, online experiments (observations), offline experiments (research on acquired images), live broadcasting, social network + community, and support. 2 objectives and benefits of gloria project the world-wide network of robotic telescopes (gloria) does not intend to compete with lsst, but our underlying idea is that “the more eyes we put on the sky, the more and greater the scientific discoveries will be achieved”. thus, in order to try to improve the way we are doing astronomy research, this project aims to build the first open access world wide network of robotic telescopes to serve citizens from around the world for free, but competing for observing time. hence gloria is indeed an “intelligent array” and it bases its intelligence in its community of users (sanchez moreno et al. 2013). like most web 2.0 projects, gloria implements a reputation-based scoring system to reward user contributions, driven by parameters such as the quality of the gathered and processed images, number of hours invested in the observations etc., as well as the votes granted by the rest of the community that finally evaluate the quality of the work done. at the end of month 24 (october 2013) the consortium produced a standard for adding new telescopes and experiments to this network. unlike other private, profitable ongoing initiatives, the gloria network will provide a free, twofold service to the community: 1. giving citizens (including professional and amateur astronomers) free access to a network of 17 robotic telescopes spread into 4 continents and both hemispheres. in the case of professional astronomers, preferences will be given to those of developing countries who will lack astronomical facilities in their own nations. 2. giving citizens an easy web access to all the data collected by the robotic telescopes. beside giving service to research, in order to manage gloria main objective of fostering astronomical research, by allowing near continuous observations of a given target thanks to its world-wide network facilities, the following additional tangible outcomes will be pursued: 1. gloria will enable more telescopes to join, producing methodology, standards, software, and documentation oriented to teach professional and amateurs astronomers to robotize their telescopes and to integrate them into the gloria network. 2. gloria will enable more research goals to be pursued, producing methodology, standards, software, and documentation oriented to teach amateurs and professional astronomers to design new web experiments and to integrate them into the gloria network. 3. gloria will encourage participation in order to increase in number its community of citizen scientists. newcomers are very welcomed. 4. gloria will give free access to knowledge to everybody. all knowledge (software, manual, standards, documents, astronomical images, data, etc) produced by the gloria consortium and by the community will have copyleft licenses. 5. gloria intends to continue in the future. since the maintenance of the distributed telescopes is paid by their owners, the cost of the core part is very cheap. the consortium believes that an economic model based in public and private subventions and donations is very possible for a project like this to go on beyond 2014. 2.1 impact the gloria partners really believe in the enormous power of astronomy as a center of interest in scientific and human training of our young people. the children of today will be the astronomers of tomorrow. in order to enroll newcomers and awaken interest in astronomy among children, during 2012-14 we will organize the live internet broadcast of 5 astronomical events: 4 eclipses and a transit of venus, which will be made from the gloria network, with associated activities in all schools of the partners countries, with the aim of getting students and teachers participating in researchbased science education and improving their motivation to push the barriers of science education further. 2.2 the gloria consortium gloria brings together a critical mass of scientific and technological leaders with extensible expertise in tic and astronomical research, dissemination and developed of technological projects. the gloria consortium includes these institutions: upm–technical university of madrid computer science faculty–spain (project coordination); auav–astronomical institute, academy of sciences of the czech republic; csic–spanish research council–spain (scientific coordination); cvtu– czech technical university–czech republic, ip-ascr– institute of physics of the academy of sciences of the czech republic; iac–astrophysic institute of canarias–el teide observatory–spain; inaf–istituto nazionale di astrofisica–italy; sao–special astrophysical observatory–russia; ucdnuid–university college dublin–ireland; uc–university of chile– chile; uma–university of málaga–spain; uoxf– oxford university–u.k.; uniwarsaw–university of warsaw–poland. 304 gloria and study of cataclysmic variables 3 gloria and cataclysmic variables cataclysmic variables (cvs) represent important targets for gloria observations mainly because of the unique properties of their activity. the category of these systems as a whole displays a very broad range of the optical activity observable by the common ccd photometers (e.g. flickering on the timescale of minutes and the amplitude of several tenths of magnitude, the orbital modulation sometimes with the amplitude of more than a magnitude, outbursts or transitions between the high and low states on the timescales of days or weeks and the amplitudes of several magnitudes). the basic properties of the optical emission of cvs need to be taken into account for the optimal strategy of monitoring of these targets in citizen science mode. we will show some examples of activity of various types of cvs which can be the targets for the gloria observers. 3.1 distribution of cataclysmic variables the sky generally, celestial objects are not distributed uniformly in the sky mainly because of the structure of our galaxy. as shown by warner (1995), most known cvs are observed over the whole sky. the reason is that cvs represent rather low-luminosity population, hence only the nearby such objects are bright enough to be observed. this relatively isotropic distribution implies that all the observatories at any latitude, included in gloria, will have the opportunity to choose a sufficiently large amount of suitable cvs. only known classical novae concentrate toward the galactic plane and the galactic bulge. the reason for this is the high optical luminosity in their explosion, hence even the objects at the large distances from the earth are observable. the direction toward the center of the galaxy is therefore also suitable for searching for explosions of classical novae with a monitor. this favors the gloria observatories located in the southern hemisphere. 3.2 peak magnitudes of cvs the catalog of downes et al. (2001) shows that the apparent peak magnitudes of various cv types differ from each other. the reason is that there is a significantly different typical optical luminosity of each type. the explosions of classical novae dominate in the brightest part of the distribution. their peak magnitude is often near 9–10 (but some of them may be even brighter), so they are easily detectable even by small-aperture, wide-field monitors. the brightest known dwarf novae in outburst can reach about 10–12 mag but about 13 mag is more typical. cvs with strongly magnetized white dwarf (polars) (e.g. warner 1995) are rather faint systems (13–20 mag), so their observing requires the largest available telescopes. 3.3 prospects of cv investigation by gloria below we give some examples of cv types suitable for investigation within gloria project. dwarf novae with their large-amplitude (3–5 mag) outbursts with the typical recurrence time from weeks to months (e.g. warner 1995; hameury et al. 1998) are definitely very promising targets even for the modest telescopes included in gloria. we also emphasize the need of observing the outbursts of some intermediate polars (i.e. cvs with a mildly magnetized white dwarf) (e.g. warner 1995). these systems may be promising gloria targets, as illustrated by the optical activity of do dra/2a 1150+720 (fig. 1). what typical features of activity of this object can be explored by gloria telescopes? (1) rare strong outbursts with the amplitude of 4–5 mag, lasting for several days and separated by ∼800 days. these short outbursts with the decay faster than exponential and faster than predicted for non-magnetic dwarf novae are in accordance with angelini & verbunt’s (1989) model for outbursts in a disk with a large missing inner region (caused by the magnetosphere of the white dwarf). (2) strong and unpredictable fluctuations on a timescale of several days or weeks (much longer than the orbital period). figure 1: the light curve of the intermediate polar do dra with rare outbursts. the points denote the one-day means, the v symbols represent the upper limits of brightness. adapted from šimon (2000). the short outbursts or rather flares of uncertain origin may occur in various intermediate polars but since they are short, extensive monitoring is needed to detect them (van amerongen & van paradijs 1989; warner 1995). supersoft x-ray sources (van den heuvel et al. 1992) represent another example of interesting objects to be investigated by gloria experiments. since the very soft x-ray emission is easily absorbed both inside 305 r. hudec, v. šimon the source and the interstellar medium, only a small fraction of supersoft x-ray sources is observable in the x-ray spectral region. the specific spectroscopic properties of their optical emission give a hope to discover more such objects even without detecting their x-ray emission (steiner & diaz 1998). the system v sge (herbig et al. 1965) can be considered to be a member of this category (steiner & diaz 1998; greiner & van teeseling 1998). since it is relative bright (10– 13 mag(v )), it is easily observable by various robotic telescopes. v sge shows a very complicated long-term optical activity. its outbursts were observed mainly in the first half of the twentieth century (šimon & mattei (1999). however, this activity changed considerably with time. a segment of the long-term activity starting in the 1960s is displayed in fig. 2). it consists of the one-day means of brightness. the high/low state transitions occurred in some time segments. they were separated by the so-called flat segments (e.g. in the surroundings of jd 2 445 000), in which only small fluctuations of brightness remained. we can thus conclude that observing at various times captures this object in different phases of its strong activity. figure 2: long-term activity of the supersoft xray source v sge in the optical band (aavso and afoev data). notice the large variations of activity on the timescale of several years (from the segments with only small fluctuations to the segments with the large-amplitude transitions between the high and low states). adapted from šimon & mattei (1999). dividing the series of the optical (ccd and photoelectric) observations of v sge, obtained at different nights (fig. 3), into the groups according to the states of the long-term activity enables to obtain the profile of the typical orbital modulation (the orbital ephemeris of smak 1995 was used). this figure shows the intensity transfered from magnitudes. although the profile of the orbital modulation can be resolved in the individual night series, the orbital period of 12.34 hours (smak 1995) does not enable to observe the whole period from a given observing site, so a multisite campaign is desirable, especially because the orbital modulation and the long-term variations of brightness are superimposed. this means that observations obtained by gloria may provide valuable information. when the intensity of a given nigh series belonging to a given such group is shifted to match a template, it is possible to obtain the mean profile of the orbital modulation characteristic for the group (fig. 3). figure 3: photoelectric and ccd observations of v sge in the intensity scale, folded with the orbital phase (the orbital ephemeris of smak 1995). the sorting into three groups is marked. adapted from šimon et al. (2002). the cv in pegasus (found in eruption in may 8, 2010) can serve as an excellent example of a bright object which can be easily detected by amateurs and small cameras and/or telescopes. this object represented a rare type of dwarf nova with very long recurrence time (67 years) but with very large amplitudes (hudec 2010). we note that the second flare of this cv in 1942 was found using astronomical photographic archives which are perfectly suitable for this kind of work, i.e. searches for past historical flare events (hudec 1999 and 2007, hudec et al. 2012). only very few such objects are known because they can easily escape detection. in this regard, we also note that the sky contains many yet unknown transients of various kinds. they are identified and attract attention only when they show themselves in their outbursts. the follow-up observations then can lead to the study of their type. one such object detected by a very small telescope is the lǐska flare star. it is an example of the astronomical discovery of a very rare event with a small amateur telescope (lǐska et al. 2014). we suggest that this object is a new flare dme star of uv ceti type based on its color, x-ray emission, and the properties of the optical light curve of the flare, identical with the x-ray sources rx j1118.3+1347 and 2e 1115.8+1403.in simbad, these 2 sources are given as different sources but this is probably wrong and it is only one source. the object may be similar to the exosat x-ray source exo 020528+1454.8 = g 035–027 with a dme flare star counterpart detected by hudec et al. (1988). 306 gloria and study of cataclysmic variables figure 4: photoelectric and ccd observations of v sge from fig. 3, now sorted into three groups according to their mean intensity (the orbital ephemeris of smak 1995). the individual curves of each group were shifted in intensity to match the template. the scale of the ordinate is identical for all three plots. adapted from šimon et al. (2002). the observations and research of cvs are expected to be supervised by gloria scientists. everybody can join us at http://gloria-project.eu, and this involves both citizen science and students, as well as professional astronomers. typically 20 percent of robotic telescopes (rts) observing time is expected to be used for gloria users, but some telescopes such as bart will provide up to 40 percent. acknowledgement gloria is project supported by the eu fp7 program (no. 283783). the scientific part of the study is linked to the the ga cr grant 13-39464j (optical analyses) while the long–term x–ray analyses are related to ga cr grant 113-33324s. we used the observations from the aavso international database (massachusetts, usa) and the afoev database operated in strasbourg, france. we thank the variable star observers worldwide whose observations contributed to this analysis as well. references [1] angelini, l., verbunt, f., 1989, mnras, 238, 697 doi:10.1093/mnras/238.3.697 [2] downes, r. a., et al., 2001, pasp, 113, 764 doi:10.1086/320802 [3] greiner, j., van teeseling, a., 1998, a&a, 339, l21 [4] hameury, j.-m., et al., 1998, mnras, 298, 1048 doi:10.1111/j.1365-8711.1998.01773.x [5] herbig, g. h., et al., 1965, apj, 141, 617 doi:10.1086/148149 [6] hudec, r., 2010, the astronomer’s telegram, no. 2619 [7] hudec, r., et al., 1988, astronomical institutes of czechoslovakia, bulletin (issn 0004-6248), vol. 39, no. 5, p. 296-302 [8] hudec, r., et al., 2012, acta polytechnica, ibws2011 proceedings, 1(52) [9] hudec, r., 2007, in exploring the cosmic frontier: astrophysical instruments for the 21st century. eso astrophysics symposia, isbn 978-3-540-39755-7. germany, p.79 doi:10.1007/978-3-540-39756-4 25 [10] hudec, r., 1999, an introduction to the world’s large plate archives, acta historica astronomiae, vol.6, 28 [11] lǐska, j., et al., 2014, acta polytechnica, accepted [12] sanchez moreno, f. m., et al., 2013, glorora documentation package [13] šimon, v., mattei, j. a., 1999, a&as, 139, 75 [14] šimon, v., 2000, a&a, 360, 627 [15] šimon, v., hric, l., petŕık, k., et al., 2002, a&a, 393, 921 [16] smak, j., 1995, acta astron., vol. 45, 361 [17] steiner, j. e., diaz, m. p., 1998, pasp, 110, 276 doi:10.1086/316139 [18] van amerongen, s., van paradijs, j., 1989, a&a, 219, 195 [19] van den heuvel, e. p. j., et al., 1992, a&a, 262, 97 [20] warner, b., 1995, cataclysmic variable stars, cambridge univ. press, cambridge doi:10.1017/cbo9780511586491 307 http://dx.doi.org/10.1093/mnras/238.3.697 http://dx.doi.org/10.1086/320802 http://dx.doi.org/10.1111/j.1365-8711.1998.01773.x http://dx.doi.org/10.1086/148149 http://dx.doi.org/10.1007/978-3-540-39756-4_25 http://dx.doi.org/10.1086/316139 http://dx.doi.org/10.1017/cbo9780511586491 introduction objectives and benefits of gloria project impact the gloria consortium gloria and cataclysmic variables distribution of cataclysmic variables the sky peak magnitudes of cvs prospects of cv investigation by gloria 139 acta polytechnica ctu proceedings 1(1): 139–145, 2014 139 doi: 10.14311/app.2014.01.0139 latest results from the fermi gamma-ray telescope aldo morselli1 1infn roma tor vergata corresponding author: aldo.morselli@roma2.infn.it abstract can we learn about new physics with astronomical and astro-particle data? since its launch in 2008, the large area telescope, onboard of the fermi gamma-ray space telescope, has detected the largest amount of gamma rays in the 20 mev 300 gev energy range and electrons + positrons in the 7 gev1 tev range, opening a new observational window on a wide variety of astrophysical objects. keywords: gamma ray gamma ray detectors dark matter. 1 introduction the fermi observatory carries two instruments onboard: the gamma-ray burst monitor (gbm) [1] and the large area telescope (lat) [2]. the lat is a pair conversion telescope for photons above 20 mev up to a few hundreds of gev. the field of view is ∼2.4 sr and lat observes the entire sky every ∼ 3 hours (2 orbits). these features make the lat a great instrument for dark matter (dm) searches. the operation of the instrument through the first three years of the mission was smooth at a level which is probably beyond the more optimistic prelaunch expectations. the lat has been collecting science data for more than 99% of the time spent outside the south atlantic anomaly (saa). the remaining tiny fractional down-time accounts for both hardware issues and detector calibrations [4], [5]. more than 650 million gamma-ray candidates (i.e. events passing the background rejection selection) were made public and distributed to the community through the fermi science support center (fssc) 1. over the first three years of mission the lat collaboration has put a considerable effort toward a better understanding of the instrument and of the environment in which it operates. in addition to that, a continuous effort was made to in order to make the advances public as soon as possible. in august 2011 the first new event classification (pass 7) since launch was released, along with the corresponding instrument response functions (and a release of a new event class ’pass 7 reprocessed’ is planned for the near future). compared with the prelaunch (pass 6) classification, it features a greater and more uniform exposure, with a significance enhancement in acceptance below 100 mev. the fermi lat results on the extragalactic sky will be covered by benoit lott [3]. here we will present the main results regarding the indirect dark matter searches and the origin of cosmic rays. 2 indirect dark matter searches one of the major open issues in our understanding of the universe is the existence of an extremely-weakly interacting form of matter, the dark matter, supported by a wide range of observations including large scale structures, the cosmic microwave background and the isotopic abundances resulting from the primordial nucleosynthesis. complementary to direct searches being carried out in underground facilities and at accelerators, the indirect search for dm is one of the main items in the broad fermi science menu. the word indirect denotes here the search for signatures of weakly interactive massive particle (wimp) annihilation or decay processes through the final products (gammarays, electrons and positrons, antiprotons) of such processes. among many other ground-based and spaceborne instruments, the lat plays a prominent role in this search through a variety of distinct search targets: gamma-ray lines, galactic and isotropic diffuse gamma-ray emission, dwarf satellites, cr electrons and positrons. 2.1 galactic center the galactic center (gc) is expected to be the strongest source of γ-rays from dm annihilation, due to its coincidence with the cusped part of the dm halo density profile [10], [11], [12]. a preliminary analysis of the data, taken during the first 11 months of the fermi satellite operations is presented in [13], [14]. 1the fssc is available at http://fermi.gsfc.nasa.gov/ssc 139 http://dx.doi.org/10.14311/app.2014.01.0139 aldo morselli figure 1: derived 95% c.l. upper limits on wimp annihilation cross sections in the milky way halo, for the muon (left) and tau (right) annihilation channels. the purple and blue contours show pamela and fermi positron excess dm interpretation constraint regions. the diffuse gamma-ray backgrounds and discrete sources, as we know them today, can account for the large majority of the detected gamma-ray emission from the galactic center. nevertheless a residual emission is left, not accounted for by the above models [13], [14]. improved modeling of the galactic diffuse model as well as the potential contribution from other astrophysical sources (for instance unresolved point sources) could provide a better description of the data. analyses are underway to investigate these possibilities. 2.2 galactic halo in order to minimize uncertainties connected with the region of the galactic center, analysis [15] considered a region of interest consisting of two off-plane rectangles (5◦ ≤ |b| ≤ 15◦ and |l| ≤ 80◦) and searched for continuum emission from dark matter annihilation or decay in the smooth galactic dark matter halo. they considered two approaches: a more conservative one in which limits were set on dm models assuming that all gamma ray emission in that region might come from dark matter (i.e. no astrophysical signal is modeled and subtracted). in a second approach, dark matter source and astrophysical emission was fit simultaneously to the data, marginalizing over several relevant parameters of the astrophysical emission. as no robust signal of dm emission is found, dm limits are set. these limits are particularly strong on leptonic dm channels, which are hard to constrain in most other probes (notably in the analysis of the dwarf galaxies, described below). this analysis strongly challenges dm interpretation [16] of the positron rise, observed by pamela [17] and fermi lat [18, 19] (see figure 1). 2.3 dwarf galaxies dwarf satellites of the milky way are among the cleanest targets for indirect dark matter searches in gamma-rays. they are systems with a very large mass/luminosity ratio (i.e. systems which are largely dm dominated). the lat detected no significant emission from any of such systems and the upper limits on the γ-ray flux allowed us to put very stringent constraints on the parameter space of well motivated wimp models [20]. a combined likelihood analysis of the 10 most promising dwarf galaxies, based on 24 months of data and pushing the limits below the thermal wimp cross section for low dm masses (below a few tens of gev), has been recently performed [21]. the main advantages of the combined likelihood are that the analysis can be individually optimized and that combined limits are more robust under individual background fluctuations and under individual astrophysical modelling uncertainties than individual limits. the derived 95% c.l. upper limits on wimp annihilation cross sections for different channels are shown in figure 2 (top). the most generic cross section (∼ 3 · 10−26cm3s−1 for a purely s-wave cross section) is plotted as a reference. these results are obtained for nfw profiles [22] but for cored dark matter profiles the j-factors for most of the dsphs would either increase or not change much so these results includes j-factor uncertainties [21]. with the present data we are able to rule out large parts of the parameter space where the thermal relic density is below the observed cosmological dark matter density and wimps are dominantly produced non-thermally, e.g. in models where supersymmetry breaking occurs via anomaly mediation for the mssm model [20]. future improvements (apart from increased amount of data) will include an improved event selection with a larger effective area and photon energy range, and the inclusion of more satellite galaxies. in figure 2 (bottom) are shown the predicted upper limits in the hypothesis of 10 years of data instead of 2; 30 dsphs in140 latest results from the fermi gamma-ray telescope stead of ten (supposing that the new optical surveys will find new dsph); spatial extension analysis (source extension increases the signal region at high energy e ≥ 10 gev,m ≥ 200 gev). figure 2: derived 95% c.l. upper limits on wimp annihilation cross sections for different channels. down: predicted 95% c.l. upper limits on wimp annihilation cross sections in 10 years for bbar channel. other complementary limits were obtained with the search of possible anisotropies generated by the dm halo substructures [23], the search for dark matter satellites [24] and a search for high-energy cosmic-ray electrons from the sun [25]. 2.4 gamma-ray lines a line at the wimp mass, due to the 2γ production channel, could be observed as a feature in the astrophysical source spectrum [12]. such an observation would be a “smoking gun” for wimp dm as it is difficult to explain by a process other than wimp annihilation or decay and the presence of a feature due to annihilation into γz in addition would be even more convincing. no significant evidence of gamma-ray line(s) has been found in the first two years of data from 7 to 200 gev [26] (see also [27]). recently, the claim of an indication of line emission in fermi-lat data [28, 29] has drawn considerable attention. using an analysis technique similar to [27], but doubling the amount of data as well as optimizing the region of interest for signal over square-root of background, [28] found a (trial corrected) 3.2 σ significant excess at a mass of ∼ 130 gev that, if interpreted as a signal would amount to a cross-section of about < σv >∼ 10−27cm3s−1. figure 3: dark matter annihilation 95% cl cross section upper limits into γγ for the einasto profile for a circular region of interest (roi) with a radius rgc = 16 ◦ centered on the gc with |b| < 5◦ and |l| > 6◦ masked. the signal is found to be concentrated on the galactic centre with a spatial distribution consistent with an einasto profile [30]. this is marginally compatible with the upper limit presented in [26]. in the analysis of the 4 year data the fermi lat team has improved over the two year paper in three important aspects: i) the search was performed in five regions of interest optimized for dm search under five different assumptions on the morphology of the dm signal, ii) new improved data set (pass 7 reprocessed) was used, as it corrects for loss in calorimeter light yield due to radiation damage during the four years of the fermi mission and iii) the energy dispersion was improved by adding a 2nd dimension to the previously used triple gaussian probability distribution function (pdf) model, leading to a so called ’2d’ pdf (such procedure is shown to increase the sensitivity to a line detection by 15%).. in that analysis [31] no globally significant lines have been fond and new limits to this dm annihilation channel were set (see figure 3). in a close inspection of the 130 gev feature it was found that indeed there exist a 135 gev signal at 4.01σ local significance, when a ’1d’ point spread 141 aldo morselli function (psf) and old data sets were used (consistently with what [28, 29] have found). however, the significance drops to 3.35σ (local, or ≤ 2σ global significance once trials factors are taken into account). in addition, a weaker signal is found at the same energy in the control sample (in the earth limb), which might point to a systematics effectpresent in this data set. in order to examine this possibility weekly observations of the limb are scheduled, and a better understanding of a nature of the excess in the control sample should be available soon. a new version of the event-level reconstruction and analysis framework (called pass 8 ) is foreseen soon from the fermi lat collaboration. with this new analysis software we should increase the efficiency of the instrument at high energy and have a data set based on independent event analysis thus gaining a better control of the systematic effects. 2.5 the cosmic ray electron spectrum the experimental information available on the cosmic ray electron (cre) spectrum has been dramatically expanded with a high precision measurement of the electron spectrum from 7 gev to 1 tev by the fermi lat [18], [19]. the spectrum shows no prominent spectral features and it is significantly harder than that inferred from several previous experiments. recently the fermi-lat collaboration performed a direct measurement of the absolute e+ and e− spectra, and of their fraction [35]. as the fermi-lat does not carry a magnet, analysis took advantage of the fact that due to its magnetic field, the earth casts a shadow in electron or positron fluxes in precisely determined regions. as a result, this measurement confirmed a rise of the positron fraction observed by pamela, between 20 and 100 gev and determined for the first time that it continues to rise between 100 and 200 gev (see figure 4). these measurements show that a new component of e+ and e− are needed with a peak at ∼ 1 tev. the temptation to claim the discovery of dark matter from detection of electrons and positrons from annihilation of dark matter particles is strong but there are competing astrophysical sources, such as pulsars, that can give a strong flux of primary positrons and electrons (see [16] and references therein). at energies between 100 gev and 1 tev the electron flux reaching the earth may be the sum of an almost homogeneous and isotropic component produced by galactic supernova remnants and the local contribution of a few pulsars with the latter expected to contribute more and more significantly as the energy increases. if a single nearby pulsar gives the dominant contribution to the extra component a large anisotropy and a small bumpiness should be expected; if several pulsars contribute the opposite scenario is expected. so far no positive detection of cre anisotropy was reported by the fermi-lat collaboration, but some stringent upper limits were published [34] and the pulsar scenario is still compatible with these upper limits. energy (gev) 1 10 210 p o si tr o n f ra ct io n −110 fermi 2011 pamela 2009 ams 2007 heat 2004 figure 4: positron fraction measured by the fermi lat and by other experiments [32, 33, 17]. the fermi statistical uncertainty is shown with error bars and the total (statistical plus systematic uncertainty) is shown as a shaded band. after the conference the ams-02 collaboration presented the result on the positron fraction [36] that confim the positron ratio rise observed by pamela and fermi and extend it up to 350 gev. forthcoming measurements from ams-02 and calet are expected to reduce drastically the uncertainties on the propagation parameters by providing more accurate measurements of the spectra of the nuclear components of cr. fermi-lat and those experiments are also expected to provide more accurate measurements of the cre spectrum and anisotropy looking for features which may give a clue of the nature of the extra component. 3 origin of cosmic rays cosmic rays are particles (mostly protons) accelerated to relativistic speeds. despite wide agreement that supernova remnants (snrs) are the sources of galactic cosmic rays, unequivocal evidence for the acceleration of protons in these objects is still lacking. when accelerated protons encounter interstellar material they produce neutral pions, which in turn decay into gamma rays. this offers a compelling way to detect the acceleration sites of protons. the identification of pion-decay gamma rays has been difficult because high-energy electrons also produce gamma rays via bremsstrahlung and inverse compton scattering. we detected the characteristic pion-decay feature in the gamma-ray spectra of two snrs, ic 443 and w44, with the fermi large area 142 latest results from the fermi gamma-ray telescope telescope. this detection provides direct evidence that cosmic-ray protons are accelerated in snrs. energy (ev) 810 910 1010 1110 1210 ) -1 s -2 d n /d e ( e rg c m 2 e -1210 -1110 -1010 best-fit broken power law fermi-lat veritas (acciari et al. 2009) magic (albert et al. 2008) agile (tavani et al. 2010) -decay0π bremsstrahlung bremsstrahlung with break ic 443 energy (ev) 810 910 1010 1110 1210 ) -1 s -2 d n /d e ( e rg c m 2 e -1210 -1110 -1010 best-fit broken power law fermi-lat agile (giuliani et al. 2011) -decay0π bremsstrahlung bremsstrahlung with break w44 figure 5: (a and b) gamma-ray spectra of ic 443 (a) and w44 (b) as measured with the fermi-lat [37] . color-shaded areas bound by dashed lines denote the best-fit broadband smooth broken power law (60 mev to 2 gev), gray-shaded bands show systematic errors below 2 gev due mainly to imperfect modeling of the galactic diffuse emission. at the high-energy end, tev spectral data points for ic 443 from magic [38] and veritas [39] are shown. solid lines denote the best-fit pion-decay gamma-ray spectra, dashed lines denote the best-fit bremsstrahlung spectra, and dash-dotted lines denote the best-fit bremsstrahlung spectra when including an ad hoc low-energy break at 300 mev c−1 in the electron spectrum. these fits were done to the fermi lat data alone (not taking the tev data points into account). magenta stars denote measurements from the agile satellite for these two snrs, taken from [40] and [41], respectively. figure 5 shows the spectral energy distribution obtained for ic 443 and w44 through maximum likelihood estimation. the normalizations of the fluxes of ic 443 and w44 and those of neighboring sources and of the galactic diffuse model, were left free in the fit for each bin. to determine whether the spectral shape could indeed be modeled with accelerated protons, we fit the lat spectral points with a π0-decay spectral model, which was numerically calculated from a parameterized energy distribution of relativistic protons. the measured gamma-ray spectra, in particular the low-energy parts, matched the π0-decay model (fig.5) [37]. the π0-decay gamma rays are likely emitted through interactions between “crushed cloud” gas and relativistic protons, both of which are highly compressed by radiative shocks driven into molecular clouds that are overtaken by the blast wave of the snr unless we introduce in an ad hoc way an additional abrupt break in the electron spectrum at 300 mev c−1 (fig.5 dash-dotted lines), the bremsstrahlung models do not fit the observed gamma-ray spectra. if we assume that the same electrons are responsible for the observed synchrotron radiation in the radio band, a lowenergy break is not expected to be very strong in the radio spectrum and thus the existing data do not rule out this scenario. the introduction of the low-energy break introduces additional complexity and therefore a bremsstrahlung origin is not preferred. finding evidence for the acceleration of protons has long been a key issue in attempts to elucidate the origin of cosmic rays. our spectral measurements down to 60 mev enable the identification of the π0-decay feature, thus providing direct evidence for the acceleration of protons in snrs. the proton momentum distributions, well-constrained by the observed gamma-ray spectra, are yet to be understood in terms of acceleration and escape processes of high-energy particles. 4 conclusions fermi turned four years in orbit on june, 2012, and it is definitely living up to its expectations in terms of scientific results delivered to the community. the mission is planned to continue at least four more years (likely more) with many remaining opportunities for discoveries. acknowledgement the fermi lat collaboration acknowledges support from a number of agencies and institutes for both development and the operation of the lat as well as scientific data analysis. these include nasa and doe in the united states, cea/irfu and in2p3/cnrs in france, asi and infn in italy, mext, kek, and jaxa in 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[40] m. tavani, et al., [agile coll.], apj 710, l151 (2010). doi:10.1103/physrevlett.110.141102 [41] a. giuliani, et al., [agile coll.], apj 742, l30 (2011). doi:10.1126/science.1231160 145 http://dx.doi.org/10.1103/physrevd.82.092003 http://dx.doi.org/10.1103/physrevlett.110.141102 http://dx.doi.org/10.1126/science.1231160 introduction indirect dark matter searches galactic center galactic halo dwarf galaxies gamma-ray lines the cosmic ray electron spectrum origin of cosmic rays conclusions 3 acta polytechnica ctu proceedings 2(1): 3–20, 2015 3 doi: 10.14311/app.2015.02.0003 the golden age of cataclysmic variables and related objects (old and news) f. giovannelli1, l. sabau-graziati2 1inaf istituto di astrofisica e planetologia spaziali, area di ricerca di tor vergata, via fosso del cavaliere, 100 i00177 roma, italy 2intadpt. cargas utiles y ciencias del espacio, c/ra de ajalvir, km 4 e28850 torrejón de ardoz, madrid, spain corresponding author: franco.giovannelli@iaps.inaf.it abstract in this paper we review cataclysmic variables (cvs) discussing several hot points about the renewing interest of today astrophysics about these sources. we will briefly discuss also about classical and recurrent novae, as well as the intriguing problem of progenitors of the type ia supernovae. this paper is an extended and updated version of the review by giovannelli (2008). because of limited length of the paper and our knowledge, this review does not pretend to be complete. however, we would like to demonstrate that the improvement on knowledge of the physics of our universe is strictly related also with the multifrequency behaviour of cvs, which apparently in the recent past lost to have a leading position in modern astrophysics. keywords: cataclysmic variables dwarf novae intermediate polars polars novae optical spectroscopy photometry sub-mm ir radio uv x-rays multifrequency astrophysics x-ray binary systems high energy astrophysics. 1 introduction in the 1950s it was recognized that the various phenomena displayed by the cvs are all the consequence of accretion of matter onto a white dwarf (wd) from a low mass donor star (e.g., warner, 1976; 1995a). cvs are binary systems in which the primary component is a wd (mwd ∼ 1 m�) and the secondary is a late type main sequence star (ms ≤ 1 m�) (e.g., smak, 1985a). mass transfer is strongly depending, besides the orbital parameter of the system, on the magnetic field intensity at the surface of the primary. such process produces a large fan of behaviour that are detectable in different energy ranges: from radio to x-rays, and even in γ-rays. the orbital periods of cvs are ranging from ∼ 80 m to ∼ 12 h with a distribution showing a gap between 2 and 3 hours, in which few systems have been detected. in the past this gap was empty and this was the reason because was nicknamed ’period gap’. more than 600 cvs are known, most of them discovered through optical observations, and some, especially those in which the magnetic field of the wd is strong, discovered through x-ray observations, but with the detectors of the second and further generations, since cvs are in general not very bright in x-ray energy range. those with known or suspected orbital period are listed by ritter & kolb (1998). the first cv detected in the x-ray range, with rocket experiments, was the dwarf nova ss cyg (rappaport et al., 1974; heise et al., 1978). the uhuru satellite detected two cvs, which were not recognized as such. warner (1976) proposed the identification of 4u 1249-28 with ex hya, and the variable am her, which on further optical studies was recognized as a cv (forman et al., 1978). the magnetic field in these two systems is strong (≈ 107–108 g). a few dozen cvs were detected in x-rays with heao-1 satellite, with exosat, and with the einstein satellite (e.g., reviews of cordova & mason, 1983; cordova, 1995). later verbunt et al. (1997) recognized 91 cvs from a sample of 162 systems with known or suspected binary periods by using data of the rosat xrt-pspc all sky survey. historically, because cvs were observed photometrically and without seeming to follow any regular pattern, they were named with the term cataclysmic (from the greek word kataklysmos = flood, storm; hack & la dous, 1993). as collecting of observational data progressed it became apparent that these objects were regular binary systems which for some reason changed in brightness; some of them also regularly (recurrent novae and dwarf novae) while some others only once (classical novae). therefore the classification of cvs was based on the optical outburst properties, by which one may distinguish four groups of cvs: (i) classical novae; (ii) recurrent novae; (iii) dwarf novae; (iv) nova3 http://dx.doi.org/10.14311/app.2015.02.0003 f. giovannelli, l. sabau-graziati like objects (e.g., giovannelli & martinez-pais, 1991 and references therein; ritter, 1992; giovannelli, 2008). this classification, however, is neither self-consistent nor adequate and it is much better to consider primarily the observed accretion behaviour (smak 1985b). one obvious advantage of such an approach is connected with the time scales of various accretion phenomena, which are sufficiently short to avoid any major observational bias: the mass accretion rates in cvs usually range from 10−11 to 10−8 m� yr −1 (patterson, 1984); the time scales are from tens of seconds (oscillations in dwarf novae at outbursts) to years (super-outbursts of su uma stars or long term variations in vy scl stars). however, in the class of nova-like objects there are two sub-classes: the dq her stars and the am her stars. in these sub-classes of cvs the wds possess magnetic fields with intensity enough high for dominating the accretion disk and all the phenomena related to it. these classes of magnetic cvs, whose names are coming from the prototypes dq her and am her took later the names of intermediate polars and polars, respectively. a short history of their discovery has been discussed by warner (1995b). fundamental papers about these subclasses are those by patterson (1994), warner (1996). the class of ips has been split into two subclasses with relatively large and relatively weak magnetic field (norton et al., 1999). one example of a system belonging to the latter subclass is do dra (previously registered as yy dra) (andronov et al., 2008). there is another class of cvs, the rare am canum venaticorum (am cvn) star systems. they have extremely short orbital periods between ∼ 10 − 65 minutes. their spectra do not show evidence for hydrogen. they appear to be helium–rich versions of cvs (e.g. warner, 1995c; nelemans, 2005). there is an old suggestion, that in these systems the mass transfer is driven by gravitational wave radiation losses, proposed by paczyński (1967), after the discovery of the prototype with an orbital period of ∼ 17 minutes (smak, 1967). depending on the magnetic field intensity at the wd, the accretion of matter from the secondary star onto the primary can occur either via an accretion disc (in the so-called non-magnetic cvs: nmcvs) or a channelling through the magnetic poles (in the case of polars: pcvs) or in an intermediate way (in the case of intermediate polars: ipcvs). cvs in a time scale of order between weeks and years flare up almost periodically, about few magnitudes in optical wavelengths; the duration of the outbursts is much shorter than the recurrence time. typical light curves for classical novae and dwarf novae of the u gem, z cam, and su uma types can be seen in ritter (1992) and e.g. in giovannelli (2008). the recurrence time-scale of outbursts in dwarf novae is correlated with their amplitude and the outburst duration is depending on the orbital period (warner, 1987). in pcvs the wd magnetic field is strong enough to make the alfvén radius greater than the circularization radius, so no accretion disc is formed and the accretion structure is fully governed by the magnetic field, which canalize the accreting matter across the field lines. owing to the intense magnetic field (∼ 10–200 mg), the wd rotation is synchronized with the binary orbital period (a few hours). however, there are few systems (v1432 aql, by cam, v1500 cyg, v4633 sgr, and cd ind) in which pspin and porb differ by around 2% or less. these are assumed to be polars that have been disturbed from synchronism by a recent nova explosion (norton, sommerscales & wynn, 2004, and the references therein). ipcv wds have moderate magnetic fields (order of a few mg); the alfvén radius is smaller than the circularization radius but it is greater than the wd radius. therefore an accretion disc is formed in these systems but being disrupted at its inner region. in ipcvs matter follows again the magnetic field lines but just inside the alfvén radius. the rotating wd is asynchronous with the binary orbital period (pspin � porb). however, there are few systems that may be best described as nearly synchronous intermediate polars (v381 vel, rxj0524+42, hs0922+1333, and v697 sco) (norton, sommerscales & wynn, 2004, and the references therein). two of these systems lie in the ‘period gap‘. probably all four systems are ips in the process of attaining synchronism and evolving into polars. the last group defined by the accretion structure criterion, nmcvs, includes those systems whose wd magnetic fields are not relevant in governing the accretion structure. in these systems the accretion disc extends down to the wd surface and a boundary layer is formed. this family shows a great diversity of observational behaviour; for this reason the historical criterion of classification is, in this case, more appropriate for distinguishing their sub-classes. however, it is simply an attempt of classification for lack of a more general physical classification (e.g., giovannelli, 1991 and references therein). indeed, in general, we can consider the wd of a cv as a gravimagnetic rotator, characterized by a mass m — accreting matter at a rate ṁ from the optical companion (the secondary star) — rotating with a velocity ~ω and having a magnetic moment ~µ, not necessarily coaxial with the rotational axis (lipunov, 1987; 1991). then the accreting system is completely characterized by the following physical parameters: mass m, accretion rate ṁ, rotational velocity ~ω, and magnetic moment ~µ. in the plane spin period of the wd (and in general of the compact object) – gravimagnetic param4 the golden age of cataclysmic variables and related objects (old and news) eter y = ṁ/µ2 it is possible to find any sort of physical conditions of gravimagnetic rotators, as discussed by e.g. giovannelli (1991). figure 1 shows the positions of few cvs in the diagram spin period of the wd – gravimagnetic parameter (in unit -42): am her (terada et al., 2010 and references therein), ae aqr (patterson, 1979; de jager et al., 1994; wynn, king & horne, 1997), dq her (patterson, 1994; zhang et al., 1995), ei uma (reimer et al., 2008), ss cyg (giovannelli & sabau-graziati 2012a). it appears evident the power of this diagram obtained by lipunov (1987) using 1 m� white dwarf. the polars am her and ae aqr lie in the zone of propeller, how they must stay, while the ipcv dq her (the prototype of this class), ei uma (a very well known ipcv), and ss cyg (whose nature as ipcv is claimed by giovannelli’s group on the base of many circumstantial proofs and also because of a cogent similarity with ei uma, e.g. giovannelli & sabau-graziati 2012a,c) lie just in the zone predicted by lipunov for such objects. figure 1: the positions of several cvs in lipunov’s diagram calculated for 1 m� white dwarf (after lipunov, 1987). as recalled by giovannelli & sabau-graziati (1999), it is evident that the properties of an outburst in cvs depend crucially on the accretion rate, the mass of the wd, and the chemical composition of its hydrogen rich envelope in which the thermonuclear runaway occurs. and the accretion process onto the wd is strongly influenced by its magnetic field intensity. indeed, the three kind of cvs (non-magnetic, polars, and intermediate polars) obey to relationships between the orbital period of the system and the spin period of the wd (warner & wickramasinghe, 1991), where the magnetic field intensity plays a fundamental role. the orbital evolution of cvs, and hence the mass-transfer rate (ṁ) from the secondary to the white dwarf is driven by magnetic braking of the secondary for long-period systems (porb > 3 hr) and gravitational radiation for shortperiod systems (porb < 2 hr). however, such a gap — which was believed true for long time — is now partially filled by the sw sex systems (e.g. rodriguez-gil, 2003; rodriguez-gil et al., 2007). the apparent ’period gap’ was due to a smaller number of systems having orbital periods in such an interval, which were escaping from the observations. therefore the investigation on the magnetic field intensities in wds is crucial in understanding the evolution of cvs systems. the fundamental parameters to be searched are the magnetic moment, the mass accretion rate and the orbital parameters of the systems. in this way it will be possible to fulfill the plane log pspin– log porb, where a priori there are not restricted ranges of magnetic moment | ~µ |, or special correlations between pspin and porb and | ~µ |. the distribution of objects in that diagram is owed to the interaction of braking torques and accretion torques, with the superposition of the observed or implied variations of the accretion rate on long time scale (> 102 yr), acting on a continuum of magnetic moments. in this way each system is completely described by those physical parameters. davis et al. (2008) applied population synthesis techniques to calculate the present day number of two types of wd-main sequence star (wdms) binaries within the ’period gap’. the first are post-common envelope binaries with secondary stars that have masses 0.17 ≤ ms/m� ≤ 0.36 (gpcebs), such that they will commence mass transfer within the period gap. the second type are systems that were cvs at some point in their past, but detached once they evolved down in orbital period to ≈ 3 h as a consequence of disrupted magnetic braking, and are crossing the ’period gap’ via gravitational radiation (dcvs). they predicted an excess of dcvs over gpcebs within the ’period gap’ of ∼ 4 to ∼ 13. this excess is revealed as a prominent peak at the location of the ’period gap’ in the orbital period distribution of the combined gpceb and dcv population. they suggest that if such a feature is observed in the orbital period distribution of an observed sample of short orbital period wdms binaries, this would strongly corroborate the disruption of magnetic braking. willems et al. (2005) and willems et al. (2007) by using population synthesis tools studied the population of nmcvs with orbital periods 1) < 2.75 h, and 2) > 2.75 h, respectively. 1) a grid of detailed binary evolutionary sequences was calculated and included in the simulations to take account of additional angular momentum losses beyond that associated with gravitational radiation and mass loss, due to nova outbursts, from the system. as a specific example, willems et al. (2005) considered the 5 f. giovannelli, l. sabau-graziati effect of a circumbinary disc to gain insight into the ingredients necessary to reproduce the observed orbital period distribution. the resulting distributions showed that the period minimum lies at about 80 minutes, with the number of systems monotonically increasing with increasing orbital period to a maximum near 90 minutes. there is no evidence for an accumulation of systems at the period minimum, which is a common feature of simulations in which only gravitational radiation losses are considered. the shift of the peak to about 90 minutes is a direct result of the inclusion of systems formed within the period gap. 2) the population of nmcvs with unevolved main– sequence–like donors at orbital periods greater than 2.75 h was investigated. in addition to the angular momentum losses associated with gravitational radiation, magnetic braking, and mass loss from the system, willems et al. (2007) also included the effects of circumbinary discs on the evolution. for a fractional mass input rate into the disc, corresponding to 3×10−4 of the mass transfer rate, the model systems exhibit a bounce at orbital periods greater than 2.75 hr. the simulations revealed that: i) some systems can exist as dwarf novae throughout their lifetime, ii) dwarf novae can evolve into novalike systems, and iii) novalike systems can evolve back into dwarf novae during their postbounce evolution to longer orbital periods. among these subclasses, novalike cataclysmic variables would be the best candidates to search for circumbinary discs at wavelengths ≥ 10 µm. the theoretical orbital period distribution is in reasonable accord with the combined population of dwarf novae and novalike systems above the period gap, suggesting the possibility that systems with unevolved donors need not detach and evolve below the period gap as in the disrupted magnetic braking model. experimental data are necessary for checking the validity of theoretical predictions. the field strength distribution of mcvs differs from that of single wds, although both cluster around 30 mg. the distribution of isolated wds extend on a wide range of magnetic field strength (∼ 105–109 g), whilst in accreting wds of cvs, as far as is presently known, there is a lack of systems at both high and low field strengths. however, the apparent absence of low field mcvs might be explained by the ips, which generally have unknown field strengths, and the lack of high field systems is still not understood (e.g., beuermann, 1998). wynn (2000) discussed the problem of accretion flows in mcvs. on the base of the ratio pspin/porb he divided the mcvs in three classes: class 1, class 2 and class 3 if such a ratio is � 0.1, ∼ 0.1, and � 0.1, respectively. for the systems in class 1 the disc equilibrium condition is clearly satisfied. those in class 2 are very unlikely to possess accretion discs. the systems in class 3 are ex hya-like systems which lie below the ’period gap’ and cannot possibly contain accretion discs. these are ex hya, ht cam, rxj1039.7-0507, and v1025 cen. these all have pspin/porb > 0.1 and porb < 2 hr. dd cir and v795 her lie within the ’period gap’ with pspin/porb ∼ 0.1 and may be included in the class 2 (norton, somerscales & wynn, 2004, and the references therein). wynn (2000) crudely classified the mcvs according to the magnetic moment and orbital period. ex hya systems have magnetic moment similar to ips above the ’period gap’ and comparable to the weakest field am her-like systems. this indicates that mcvs above the ’period gap’ will evolve to long spin periods below it. norton, wynn & somerscales (2004) investigate the rotational equilibria of mcvs. they predict that ipcvs with µ ≥ 5×1033 g cm3 and porb > 3 hr will evolve into pcvs, whilst those with µ ≤ 5 × 1033 g cm3 and porb > 3 hr will either evolve into low field strength polars that are presumably unobservable, and possibly euv emitters, or into pcvs when their fields, buried by high accretion rate, revive when the mass accretion rate reduces. warner (1996) deeply discussed torques and instabilities in ips on the base of measured spin periods of the primaries and found several important relationships between fundamental parameters of these systems, such as log ṁ vs log porb, log µ33 vs log ṁ17, log lx vs log ṁ, as shown in fig. 2 in the left, central, and right panels, respectively. there is a range of magnetic moments µ and mass transfer rates in which synchronized rotation of the primary can occur even though it possesses an accretion disc. ak et al. (2010), using available astrometric and radial velocity data, computed the space velocities of cvs with respect to the sun and investigated kinematical properties of various sub-groups of cvs. the orbital period distribution of cvs in the refined sample of 159 systems resembles that of the whole sample of cvs (e.g. connon smith, 2007). ak et al. (2010) found that the mean kinematical age (mka) of the 159 systems is mka159 = 5 ± 1 gyr. in the sample, 134 of 159 systems are non magnetic (nmcv) having mkanmcv = 4.0±1.0 gyr. in the sub–sample of nmcvs, 53 of 134 have porb < 2.62 h and their mka is 5.0±1.5 gyr, whilst 81 of 134 systems have porb > 2.62 h and their mka is 3.6±1.3 gyr. this means that cvs below the ’period gap’ are older than systems above the gap. this result in agreement with the standard evolution theory of cvs. the selection of 2.62 h as the border between the two groups of systems lies roughly in the middle of the ’period gap’, where systems have been detected. this means that the ’period gap’ does not exist anymore and the systems inside this ’gap’ are just frontier objects between systems experiencing gravitational radiation and those experiencing magnetic braking. the reason because they are not so numerous as 6 the golden age of cataclysmic variables and related objects (old and news) figure 2: left panel: intermediate polars in the plane ṁ-porb. the two lines are lines of disc stability: stable above and dwarf nova outburst below. they were computed for f=0.2 and 0.3, with rout = f × a, being a the separation of the two stars in the system (by courtesy of warner, 1996). central panel: magnetic moment in units of 1033 g cm3 versus mass accretion rate in units of 1017 g s−1. boundaries for white dwarf mass m1 = 1 m� and m1 = 0.6 m� have been computed for porb = 4 h (by courtesy of warner, 1996). right panel: mass transfer rate onto white dwarf versus (2-10 kev) x-ray luminosity (by courtesy of warner, 1996). those placed at sides could be the relative shorter time of permanence in the ’gap’, and then difficult to be detected. our opinion is that a more appropriate investigation of the class of the so-called ipcvs is necessary. indeed, such systems could show surprises if deeply studied, as for instance occurred for ss cyg. this system, usually considered a nmcv because of a classification once made by bath & van paradijs (1983), whilst since 1984 it has been claimed as an ipcv by giovannelli et al. (1985) and later confirmed several times (e.g., giovannelli & martinez-pais, 1991; giovannelli, 1996; giovannelli & sabau-graziati, 1999, gaudenzi et al., 2002). giovannelli & sabau-graziati (2012a) discussed all the circumstantial proofs in favor of the magnetic nature of ss cyg, as well as those adverse, concluding with reasonable certainty that its nature is magnetic, being the magnetic field intensity b = 1.7 ± 0.8 mg, in agreement with the value (b ≤ 1.9 mg) derived by fabbiano et al. (1981) by using x-ray, uv and optical coordinated measurements. this would teach a lesson: it is mandatory to observe cvs for long time in order to follow at least a whole period of the binary system between two successive outbursts. this is, of course, possible only for systems like dwarf novae where the almost periodical outbursts occur in time scales of weeks-months. networks of robotic telescopes can help in this matter (giovannelli & sabau-graziati, 2012b) however, we can say that cvs form a broad stellar family of highly variable and dynamical members. when it comes to explaining particulars about, e.g., the detailed interaction between the transferred matter and the wd’s atmosphere; irregularities within regular photometric behaviour; turbulent transport in the disc; or the final fate of these objects, more is missing than what is known, rendering their study ever more challenging. at least, cvs are natural multi–wavelength laboratories offering us the possibility of studying in detail the behaviour of plasma and radiation under extreme physical conditions. the understanding of stellar evolution, electromagnetism and polarization, mass and radiation transfer or 3-d geometrical effects, in a broad spectral range from hard x-rays to radio, is mandatory for improving the knowledge of the nature of cvs. variability, from milliseconds to hundreds of years, follows from different physical processes taking place in these systems and can be studied by means of several astronomical techniques. as our skills in developing further these techniques grow our understanding of the cvs insights also grows; and the more we learn about cvs the further techniques and theory develop. on the other hand, it is well known that conclusions obtained in the field of cvs have been extrapolated, upwards or downwards in scale, to other fields such as agns or lmxrbs, and vice versa. from such exchanges of information and results astrophysical research in general always benefits. rapid oscillations in cvs are particularly interesting. as reviewed by warner (2004), the rich phenomenology of dwarf nova oscillations (dnos) and quasi–periodic oscillations (qpos) observed in cvs favour the interpretation that these rapid brightness modulations (3 to 11,000 s timescales) are magnetic in nature — magnetically channelled accretion from the inner accretion disc for dnos and possible magnetically excited traveling waves in the disc for qpos. there is increasing evidence for the magnetic aspects, which extend to lower fields the well–known properties of strong field (pcvs) and intermediate strength field (ipcvs) cvs. the result is that almost all cvs show the presence of magnetic fields on their wd primaries, although for many the intrinsic field may be locally enhanced by the accretion process itself. there are many behaviour that parallel the qpos seen in x-ray binaries, with 7 f. giovannelli, l. sabau-graziati high– and low–frequency x–ray qpos resembling, respectively, the dnos and qpos in cvs. other papers about rapid oscillations in cvs are those by warner & woudt (2005) and pretorius, warner & woudt (2006). the current estimate of the space density of cvs is of ∼ 3 × 10−6 pc−3 (warner, 2001). this may be a significant underestimate of cvs space density, as discussed by patterson (1984). although densities from the most comprehensive optical palomar–green survey raises the estimate at (3−6)×10−6 pc−3, x-ray all-sky surveys give densities of ∼ 1 × 10−5 pc−3 for detected systems of low ṁ in hard x-rays (patterson, 1998). then from observational point of view, it is necessary an intensive search for the faint cvs predicted by population synthesis with orbital periods at ∼ 80 − 100 min that have passed through the orbital period minimum at ∼ 78 min and have increasing orbital periods. this research must be done among the low ṁ systems detected by x-ray surveys. thanks to its high sensitivity, integral is very useful for this purpose. up to now, it discovered several new faint cvs, with porb > 3 hr, and only one with porb < 3 hr (e.g. šimon et al., 2006; hudec et al., 2008). high speed photometry of faint cvs have shown that: i) 1 of 10, tv crv has porb = 1.509 hr (woudt & warner, 2003); ii) 5 of 13 have porb < 2 hr (woudt, warner & pretorius, 2004); iii) 1 (cal 86) of 12 has porb = 1.587 hr (woudt, warner & spark, 2005); iv) 3 of 11 have porb > 3 hr (witham et al., 2007). for reviews about cvs see the fundamental papers by robinson (1976), patterson (1984, 1994), hack & la dous (1993), and the books of warner (1995a) and hellier (2001). more recent reviews are those by connon smith (2007), and giovannelli (2008). the long review the impact of space experiments on our knowledge of the physics of the universe by giovannelli & sabaugraziati (2004) contains also a part devoted to cvs. 2 multifrequency emissions in cvs there are several components that are responsible for the total emission. deep discussions about these components can be found in the literature, but there is a recent review by giovannelli (2008, and the references therein) that exhaustively summarize their contributions to the total multifrequency emission of cvs. briefly, such components are: a) the secondary stars are cool main sequence stars with spectral type ranging from g8 to m6, corresponding to temperatures from 5,000 to 3,000 k. their contribution is mainly in red and ir regions of the electromagnetic spectrum. b) the primary stars. the temperatures of wds are known only in few cases: when they belong to high inclination systems, or when they accrete matter with a very low mass transfer rate. however, the wd temperatures range between 10,000 and 50,000 k (sion, 1986; 1991). urban & sion (2006) found that the wds in cvs above the period gap are hotter and more accretion heated (teff = 25, 793 k) than those below the gap (teff = 18, 368 k). therefore wds are expected to radiate essentially in the uv, but they can be visible also in the optical range if they are not too hot. c) the accretion disc: it does not have a homogeneous temperature, but spans a large range. since the temperature distribution in discs is poorly known, in order to obtain a rough evaluation of their contribution to the total emission it is necessary to evaluate the contributions at different frequencies of a synthetic disc constituted of black bodies at different temperatures, the temperature distribution being that of a stationary accretion disc (e.g., la dous, 1994). it then appears evident that the contribution of such an accretion disc is important in the whole range between euv and ir, depending on the choice of the disc parameters. furthermore, the uv radiation can be supplied from a zone in the vicinity of the wd (some ten stellar radii), which could contain any optically thick material left there. however, the argument of accretion discs deserves an important comment. it appeared evident that the viscosity of matter inside the accretion discs plays a fundamental role in the description of physical processes occurring there. in spite of numerous attempts in determining such a viscosity, the physical nature of that still remains largely indeterminate. the best training for study the viscosity is the subclass of dwarf novae, showing quasi-periodic outbursts which occur on a time scale from weeks to months (or even years) and are due to non-stationary accretion. meyer & meyer-hofmeister (1981, 1982, 1983) firstly discussed the physical mechanism responsible for dwarf nova outbursts which is connected with the thermal instability of the disc which occurs in the temperature range corresponding to the ionization of hydrogen. soon after smak (1984a,b) extended the study of such a mechanism. the details are summarized by smak (2002). it is important to point out that the shapes of dwarf nova light curves, which depend on a number of relevant parameters, depend also on viscosity. in particular, the characteristic time-scales observed during outbursts depend on the viscous time-scale. this provides an important and almost unique opportunity of obtaining some constraints on viscosity or – within the α disc approach (shakura & sunyaev, 1973) – of an empirical determination of α. for a more complete and detailed discussion of dwarf novae and models of their outbursts – see reviews by cannizzo (1993), osaki (1996), and lasota (2001). 8 the golden age of cataclysmic variables and related objects (old and news) d) the boundary layer: a very important zone for the emission is that of the transition between the accretion disc and the wd surface, namely the boundary layer. it is possible that all, or at least a significant fraction, of the kinetic energy of the material contained in the accretion disc must be radiated away within the geometrically very small boundary layer in order to have the possibility of the material accreting onto the wd’s surface. then, whatever the situation, one can assume the presence of a strong x-ray source at the boundary layer, which will be visible also in the euv and shortwavelength uv according to the choice of disposable theoretical parameters. most of the radiation then comes from the accretion disc and boundary layer, which contribute roughly 50% each. from the accretion disc the radiation is essentially emitted in the optical and uv, whilst from the boundary layer — optically thick (which occurs at high accretion rates) — the radiation is emitted in the soft x-ray range; when the accretion is at low rates the boundary layer is optically thin and appears as a hard thermal bremsstrahlung source. these predictions have been tested experimentally, comparing the observations of cvs in optical, uv and soft x-ray ranges (e.g. wood et al., 1989; horne et al., 1994). e) the gas stream: it is definitively optically thin and cool and contains rather little material; so, probably, its contribution to the total emission of cvs is negligible at all frequencies as source of continuum, whilst it could contribute to the formation of lines in the red and ir regions. f ) the hot line: the energy excess zone in the place where the stream comes to the disc is a shock wave. this zone was previously known as ‘the hot spot‘. its structure and radiation characteristics are still an open problem; it is visible in many systems in optical photometry (less in the ir and never in the uv) as a periodically recurring hump in the orbital light curve. its temperature must be ≤ 10, 000 k. g) hot corona or chromosphere: is a shell of optically thin and rather hot gas, below and above the accretion disc. x-ray and uv line radiation are tentatively attributed to it, whilst it probably does not contribute to the uv, optical or ir continuum emission. usually no radio emission from cvs has been measured. only upper limits for individual systems of order of a few mjy are available (e.g ≤ 10 mjy in ss cyg — cordova, mason & hjellming 1983). however, recently, körding et al. (2008) detected a radio flare from ss cyg peaked at 1.1 mjy with a duration of order 20 days, above the upper limit of 0.08 mjy. this radio flare was simultaneous with the optical long outburst peaked at about 1 jy. pringle & wade (1985) computed the contribution functions of the most important components of a cataclysmic system, previously discussed. the plot can be found also in fig. 6 of giovannelli’s paper (2008). during quiescence dwarf novae emit essentially hard x-rays (∼ 0.1–4.5 kev) and the flux distribution is rather well approximated by a thermal bremsstrahlung with ktbrems ≈ 10 kev (cordova & mason, 1983). a direct correlation between the hard x-ray/optical fluxes ratio and hβ equivalent width has been found by patterson & raymond (1985). during outburst dwarf novae emit soft x-rays (0.18– 0.5 kev) with an increase of the flux of the order of 100 or more, although most of the radiation is hidden in the euv range (cordova & mason, 1984). the soft x-ray spectra can be fitted either with black bodies at ktbb ≈ 25–30 ev or, alternatively, with bremsstrahlung spectra at ktbrems ≈ 30–40 ev. the most important features are the anti-correlation between the hard and soft x-ray emissions during the outburst cycle and the correlation between soft x-ray and optical emissions, as measured for ss cyg (watson, king & heise, 1985), or — what is the same — anti– correlation between the hard x-ray and optical emission (ricketts, king & raine, 1979). during an outburst of ss cyg there is also the correlation of optical and euv emissions, that are anti–correlated with the hard x-ray emission detected by rxte (wheatley et al., 2003). what does that mean? the uv flux and the bulk of optical flux in dwarf novae and nova-like stars originate in the accretion disc. the ir flux observed during quiescence and possibly some of the optical flux come from the secondary late-type star. the rise to an outburst either occurs simultaneously at all wavelengths when it is slow, or progressively starts later with decreasing wavelengths when it is fast, since ever more central hotter parts of the disc become involved. indeed, several dwarf novae have been observed in the uv and optical during the rise to maximum outburst brightness and their behaviour are quite similar: the uv rise lags the optical rise by up to a day (e.g., vw hyi: hassall et al. (1983); cn ori and rx and: cordova, ladd & mason (1986); wx hyi: hassall, pringle & verbunt (1985). with respect to the optical band, this lag is similar also in the euv region covered by the voyager (50–1200 å) for ss cyg (polidan & holberg, 1984) and vw hyi (polidan & holberg, 1987). this fact strongly supports the origin of the outburst being in the cooler outer part of the disc rather than in the hotter parts near the wd; therefore mininova models for the outbursts are probably excluded (cordova & howarth, 1987). the two models for triggering the outbursts, compatible with the lag observations, are then: 9 f. giovannelli, l. sabau-graziati • an instability in the secondary star which allows the transfer of more mass to the disc; • a thermal instability in the outer disc, which results in material stored there being suddenly transported through the disc. during the decline, the whole disc cools simultaneously. the contribution to the total emission from the boundary layer between the disc and the wd surface is in uv and x-ray ranges: the boundary layer is optically thin during quiescence and then emits hard x-rays, but it is optically thick during outburst and then emits soft x-rays since the radiation is thermalized before escape (la dous, 1993). from the euve (extreme ultraviolet explorer) craig et al. (1997) have shown the euv spectra of three npcvs in outburst, namely vw hyi, u gem and ss cyg. vw hyi shows the softest euv spectrum peaked at ∼ 250 å. however, its boundary layer/disc luminosity ratio — lbl/ldisc ∼ 0.2 (mauche et al., 1991) — is in contrast with the boundary layer models. in u gem the total size of the euv emitting region is comparable to that of the wd itself, which indicates that the outburst in mainly confined to the inner disc/boundary layer region (long et al., 1996). its lbl/ldisc ∼ 1. ss cyg is the hardest of the three euv nmcvs. its euv spectrum is rather complex and changed by a factor 100 during the outburst with an almost constant spectral energy distribution. quasi-coherent oscillations (∼ 7−9 s) have been detected in the euv emission (mauche, 1996). ss cyg shows no euv emission longward ∼ 130 å. mauche, raymond & mattei (1995) found that for ss cyg the relation lbl/ldisc ∼ 1, valid for the boundary layer models is strongly violated, being this ratio lbl/ldisc ∼ 0.07. and this is one more proof that ss cyg is not a nmcv — as already remarked — as claimed by giovannelli’s group. iue satellite deserves special comments since it was fundamental in improving the knowledge of cvs. a detailed review can be found in giovannelli (2008). briefly, the iue gave significant contributions on: i) knowledge of disc accreting and magnetic cvs, as extensively discussed by cordova (1995), and references therein. ii) nature of the high velocity winds. during outburst of ss cyg the spectral emission features disappear or go into absorption, some of them showing p cygni profile (e.g. civ), which clearly indicate the presence of high velocity wind from the system. the emission features appear again when the system is going into quiescence (giovannelli et al., 1990). iii) boundary layer emission. multifrequency observations show that x-ray luminosity at all outburst phases is much lower (about at least a factor 10) than the uv/optical luminosity from the disc, as expected from the models (e.g., mauche, 1998). this simply means that the boundary layer models are not correct. iv) underlying wd and its photosphere. iue provided the first evidence that the wd is heated by the dwarf nova outburst and subsequently cooled. a list of nine such systems has been reported by szkody (1998). these measurements are very difficult because of the long quiescence–outburst–quiescence cycles (from weeks to years). the short outburst period dwarf nova vw hyi cooled to 18,000 k (from 20,500 k) in the 14 days before the next outburst began (verbunt et al., 1987). v) magnetic field of the wd. indirect evaluations of magnetic field intensities in cv wds have been obtained through multifrequency observations (e.g., fabbiano et al., 1981). our feeling is that the problem of magnetic fields in wds has been underestimated in the studies of cvs. too many simplified models of disc accreting and magnetic cvs have been developed under the hypothesis that cvs can be sharply divided into three classes: polar, ip, non-magnetic. magnetic fields are smoothly varying in their intensities from one class to another. the discovery in some ips of a circularly polarized optical emission suggests that these intermediate polars will evolve into polar systems (e.g., mouchet, bonnetbidaud & de martino et al., 1998). some evidence of the continuity between the ipcvs and pcvs is coming from the detection of the sw sex systems. they have orbital periods just inside the so-called ’period gap’, which separates the two classes of ipcvs and pcvs (e.g. rodriguez-gil, 2003 and references therein; rodriguez-gil et al., 2007). looking at the homogeneous set of data coming from the iue for pcvs and ipcvs, it has been possible to obtain important information on common properties and peculiarities of these binaries (de martino, 1999 and references therein), which render the two classes rather similar in some of their uv behaviour. mouchet, bonnet-bidaud & de martino (1998), and de martino (1998) made the hypothesis that the two classes are evolutionary related. the far uv vs near uv colour–colour diagram for mcvs was constructed by de martino (1999). such a diagram was constructed measuring broad band continua in the iue short wavelength range (1420–1520 å and 1730–1830 å) and in the long wavelength range (2500–2600 å and 2850–2900 å). clearly the uv continua cannot be simply described by a single component but possess different contributions, as discussed by de martino (1999) and already noted in the past, since 1984, by giovannelli et al. (1985). araujo-betancor et al. (2005) obtained hubble space telescope (hst) stis data for a total of 11 pcvs 10 the golden age of cataclysmic variables and related objects (old and news) as part of a program aimed at compiling a homogeneous database of high–quality fuv spectra for a large number of cvs. comparing the wd temperatures of pcvs with those of nmcvs, they find that at any given orbital period the wds in pcvs are colder than those in nmcvs. the temperatures of wds in pcvs below the period gap are consistent with gravitational radiation as the only active angular momentum loss mechanism. the differences in wd effective temperatures between pcvs and nmcvs are significantly larger above the period gap, suggesting that magnetic braking in pcvs might be reduced by the strong field of the primary. araujo-betancor et al. (2005) derive a lower limit on the space density of pcvs of 1.3 × 10−6 pc−3. 3 renewed interest for cataclysmic variables before the advent of rosat x-ray satellite, mcvs were relegated to a subsection of conferences about cvs that were mainly concentrated on nmcvs. the rosat satellite discovered many mcvs that even menaced to overthrow our understanding of the secular evolution of ‘normal‘ cvs by appearing – apparently inexplicable – in the so–called ‘period gap‘ in the orbital–distribution of cvs (e.g. vrielmann & cropper, 2004). but in spite of this, cvs were not considered, in general, for many years as principal targets of high energy x-ray experiments. at the beginning of the nineties of the last century, acceleration of particles by the rotating magnetic field of the wd in intermediate polars in the propeller regime — ae aqr -– detected by ground-based cherenkov telescopes in the tev passband (e.g. meintjes et al. 1992), and tev emission from the polar am her detected by ground-based cherenkov telescopes (bhat et al. 1991) — measurements never confirmed — were the main reasons of renewed interest for cvs in the high energy astrophysicists community. the integral observatory, until the beginning of 2007, had observed over 70 percent of the sky, with a total exposure time of 40 million seconds. bird et al. (2007) published the third integral catalogue of gamma-ray sources. it contains a total of 421 gammaray objects. most have been identified as either binary stars in our galaxy containing exotic objects such as black holes and neutron stars, or active galaxies, far away in space. but a puzzling quarter of sources remain unidentified so far. they could be either star systems enshrouded in dust and gas, or cvs. integral observes in the gamma-ray band so it can see through the intervening material. it has demonstrated that it can discover sources obscured at other wavelengths. one surprise has been the efficiency with which integral has detected just one minor subclass cvs, the socalled ipcvs. initially astronomers were not sure that cvs would emit gamma rays. indeed, integral has already shown that only about one percent of them do. this fact overbearingly renewed the interest for cvs, apparently fallen into disgrace in favour of binary systems containing either neutron stars or black holes. the fourth ibis/isgri catalog reports 331 additional sources when compared to the third catalog. of these, 120 are associated with extragalactic sources, while only 25 are associated with known galactic sources, and the remainder are so far unidentied (bird et al. 2010). cvs constitute ∼ 5% of the total sources. moreover, since the cvs measured by the integral observatory are magnetic in nature, the interest for such class of objects has been addressed to evolutionary problems. the long–standing fundamental predictions of evolution theory are finally being tested observationally. all facets of the accretion process in cvs, including variability, disc winds and jets, are universal with accreting wds, neutron stars, and black holes (knigge, 2010, 2011). knigge, baraffe & patterson (2011) extensively discussed the reconstruction of the complete evolutionary path followed by cvs, based on the observed mass–radius relationship of their donor stars, following knigge (2006) that discussed the observational and theoretical constraints on the global properties of secondary stars in cvs using the semi–empirical cv donor sequence, and concluded that most cvs follow a unique evolutionary track. in the standard model of cv evolution, angular– momentum–loss (aml) below the period gap are assumed to be driven solely by gravitational radiation (gr), while amls above the gap are usually described by a magnetic bracking (mb) (rappaport, verbunt & joss (1983). knigge, baraffe & patterson (2011) with their revised model, found the optimal scale factors fgr = 2.47 below the gap and fmb = 0.66 above, whilst the standard model gives fgr = fmb = 1. this revised model describes the mass–radius data much better than the standard model. the sub-class of cvs, named classical novae (cne), which are the third more powerful stellar explosions in a galaxy, have been observed as close as a kpc and as far as galaxies in fornax cluster. the time to report on the recent renaissance in studies on cne thanks to observations with 8-10m class telescopes, high resolution spectroscopy, in synergy with observations from space carried out with swift, xmm, chandra, hst, and spitzer, coupled with recent advances in the theory of the outburst, seems now in order. moreover, the possible connection among some cv-types and sne-ia will definitively justify the renewed interest about cvs. 11 f. giovannelli, l. sabau-graziati 4 classical and recurrent novae classical novae are expected to recur on timescales from 100,000 years to just a few decades. the most important physical parameters controlling this recurrence timescale are the wd mass, and the mass accretion rate from the secondary (e.g. yaron et al. 2005). once classical nova (cn) is recorded more than once, it can be designated as “recurrent” (rn). since the wd and the binary system remain intact after an outburst, it is possible that classical novae may actually be the same as recurrent novae if observed over a long enough time period. while the interval between outbursts of recurrent novae range from 10 to 100 years, it has been estimated that the time interval for classical novae would range from about 30,000 years for a 1.3 m� wd to 100,000 years for a 0.6 m� wd. given long enough it is expected that all classical novae will be observed as recurrent novae. the long term behaviour of classical old novae, and the optical behaviour of cne in outburst were discussed by bianchini (1990), and seitter (1990), respectively. the books by cassatella & viotti (1990) and by bode & evans (2008) are very useful for studying the physics of classical novae. recurrent novae are a rare sub-class of cataclysmic variable stars; wds accreting material from a binary companion in which more than one classical nova-type outburst has been observed (see the book of hellier, 2001 for a comprehensive review of cvs). nova outbursts are suspected to be due to a thermonuclear runaway on the surface of the wd, which releases huge amounts of thermal energy once a critical pressure is reached at the base of the shell of accreted material. one of the most interesting rne is rs ophiuchi (rs oph). it is an amazingly prolific recurrent nova, with recorded outbursts in 1898, 1907, 1933, 1945, 1958, 1967, 1985 and 2006 (schaefer 2010). the short time between outbursts (∼ 20 yrs) suggests that rs oph hosts a massive wd accreting material at a significantly high rate. in the latter paper schaefer discussed not only rs oph, but also the photometric histories of all known galactic rne. classical and recurrent nova outbursts have been recently discussed by bode (2011a,b) and evans (2011). the proceedings of a conference about rs oph and recurrent phenomenon can be very useful for details (evans et al., 2008). general properties of quiescent novae have been discussed by warner (2002). the very useful book of bode & evans (2008) about classical novae examines thermonuclear processes, the evolution of nova systems, nova atmospheres and winds, the evolution of dust and molecules in novae, nova remnants, and observations of novae in other galaxies. it includes observations across the electromagnetic spectrum, from radio to gamma rays, and discusses some of the most important outstanding problems in classical nova research. of the ∼ 400 known galactic classical novae, only 10 of them are recurrent. eight of them harbour evolved secondary stars, contrary to classical novae that contain main sequence stars (darnley et al., 2011). they propose a new nova classification based on the evolutionary state of the secondary star, contrary the current schemes based on the properties of outbursts. such classification contains three groups of novae: i) main sequence nova (ms–nova); ii) sub–giant nova (sg– nova); and iii) red giant branch nova (rg–nova). rne play an important role in the studies of sn ia progenitors (surina et al., 2011). rne are likely progenitors of type–ia supernovae. in order to brave this important problem the use of archival data is the only way to answer the big question. now, huge and comprehensive set of archival rn data go back to 1890. 5 progenitors of sn ia it is well accepted by the community that type–ia sne are the result of the explosion of a carbon–oxygen wd that grows to near chandrasekhar’s limit in a close binary system (hoyle & fowler, 1960). but the debate is focussed around the different kinds of progenitors. indeed, in the past, two families of progenitor models have been proposed. they differ in the mode of wd mass increase. the first family is the so–called single degenerate (sd) model (whelan & iben, 1973), in which the wd accretes and burns hydrogen–rich material from the companion. the second family is the so–called double degenerate (dd) model, in which the merging of two wds in a close binary triggers the explosion (webbing, 1984; iben & tutukov, 1984). the two scenarios produce different delay times for the birth of the binary system to explosion. thus it is hopefully possible to discover the progenitors of type–ia sne by studying their delay time distribution (ddt). the ddt can be determined empirically from the lag between the cosmic star formation rate and type–ia sn birthrate. the energy released through runaway thermonuclear process ejects the majority of the unburnt hydrogen from the surface of the star in a shell of material moving at speeds of up to 1.5 ×103 km s−1. this produces a bright but short-lived burst of light the nova. although type–ia supernova appear to have similar origin to classical novae, there are key differences. the most important is that in a classical nova, the thermonuclear runaway occurs only on the surface of the star, allowing the wd and the binary system to remain intact (e.g. townsley & bildsten, 2005). in a type–ia 12 the golden age of cataclysmic variables and related objects (old and news) supernova, the thermonuclear runaway occurs within wd itself, completely disrupting the progenitor. this is reflected in the amount of energy released in the explosions, with classical novae releasing ∼ 1044 erg, and type–ia supernovae ∼ 1051 erg. the possible progenitors of sn ia are: i) recurrent novae; ii) symbiotic stars; iii) super-soft sources; iv) double wd binaries; and v) wds accreting material from red–giant companions. i) recurrent novae are just a subset of ordinary novae that happen to go off more than once per century. as such, they are binary systems with matter flowing off a companion star onto a wd, accumulating on its surface until the pressure gets high enough to trigger a thermonuclear runaway that is the nova. only 10 rne are known in our milky way galaxy, including: u sco (1863, 1907, 1917, 1936, 1945, 1969, 1979, 1987, 1999); t pyx (1890, 1902, 1920, 1944, 1967); t crb (1866, 1946); rs oph (1898, 1907, 1933, 1945, 1958, 1967, 1985, 2006). to recur with τrec < 100 years, rne must have: high wd mass (1.2m� < mwd < mchandra), and high accretion rate (ṁ ∼ 10−7 m� yr−1). sn ia occurs if: i) the mass ejected for each eruption is less than the mass accreted onto the wd (mejected < ṁ τrec); ii) the rate of death rne must be enough to produce the sn ia rate (rrndeath = rsnia), being rrndeath = nrn ×(0.2m�ṁ). in order to solve the problems we need to know τrec (recurrence time scale) from archive plates, nrn (number of rne in the milky way) from archive plates and aavso, ṁ (mass accretion rate onto wd) from the average in the last century, mejected (mass ejected in eruption) from pre–eruption eclipse timing. some results have been obtained for becoming optimists in solving the problem of sn ia production. indeed schaefer (2011) obtained for ci aql and u sco mejected << ṁ τrec). thus, wds are gaining mass and the latter rne will collapse as sn ia. moreover, for the milky way, m31, and lmc rrndeath ∼ nrn. then there are enough rne to supply the type–ia sn events. ii) symbiotic stars contain wds efficiently accreting material from the secondary star. in most cases they steadily burn h–rich material allowing them to grow in mass. some of these systems can produce high mass wds. in symbiotic rne (syrne) the wd mass is already very close to chandrasekhar’s limit. for instance in v 407 cyg a very massive wd is accreting material at a rate of ∼ 10−7 m� yr−1 from a mira– type companion (miko lajewska, 2011). iii) super–soft sources are probably wds that accrete material and burn hydrogen. voss & nielemans (2008) discovered an object at the position of the type– ia sn2007on in the elliptical galaxy ngc1404 on pre– supernova archival x–ray images. this result favours the accretion model (sd) for this supernova, although the host galaxy is older than the age at which the explosions are predicted in sd models. however, the dd model cannot be ruled out by this event because a hot accretion disc is probably the intermediate configuration of the system, between first wd–wd roche–lobe contact and explosion (yoon, podsiadlowski & rosswog, 2007). greggio, renzini & daddi (2008) starting from the fact that type–ia sn events occur over an extended period of time, following a distribution of delay times (ddt), discussed theoretical ddt functions that accommodate both ‘prompt‘ and ‘tardy‘ sn events derived by empirically–based ddt functions. moreover such theoretical ddt functions can account for all available observational constraints. the result is that sd/dd mix of snia’s is predicted to vary in a systematic fashion as function of cosmic time (redshift). iv) double wds binaries are systems containing two wds that can merge and giving rise to sn explosion. yoon, podsiadlowski & rosswog (2007) explored the evolution of the merger of two carbon–oxygen (co) wds. their results imply that at least some products of double co wds merger may be considered good candidates for the progenitors of type–ia sne. brown et al. (2011) and kilic et al. (2011) studied a complete colour–selected sample of double–degenerate binary systems containing extremely low mass (elm) (≤ 0.25 m�) wds. milky way disc elm wds have a merger rate of ≈ 4 × 10−5 yr−1 due to gravitational wave radiation. the elm wd systems that undergo stable mass transfer can account for about 3% of am cvn stars. the most important fact is that the elm wd systems that may detonate merge at a rate comparable to the estimate rate of underluminous sne. these sne are rare explosions estimated to produce only ∼ 0.2 m� worth of ejecta. at least 25% of elm wd sample belong to the old tick disc and halo components of our galaxy. thus, if merging elm wd systems are the progenitors of underluminous sne, transient surveys must find them in both elliptical and spiral galaxies. v) wds accreting material from red–giant companions. observations carried out by patat et al. (2008) with vlt–uves allowed to detect circumstellar material in a normal type–ia sn. the expansion velocities, densities and dimensions of the circumstellar envelope indicate that this material was ejected from the system prior to the explosion. the relatively low expansion velocities favour a progenitor system where a wd accretes material from a companion star, which is in the red–giant phase at the time of explosion. bianco et al. (2011) searched for a signature of a non–degenerate companion in three years of supernova 13 f. giovannelli, l. sabau-graziati legacy survey data. they found that a contribution from wd/red–giant binary system to type–ia sn explosions greater than 10% at 2σ, and than 20% at 3σ level is ruled out. type–ia sne are used as primary distance indicators in cosmology (e.g. phillips, 2005). phillips (2011) reviewed the near–infrared (nir) of type–ia sne concluding that such sne are essentially perfect standard candles in the nir, displaying only a slight dependence of peak luminosity on decline rate and colour. lira (1995) first noted that b–v evolution during the period from 30 to 90 days after v maximum is remarkably similar for all sn ia events, regardless of light–curve shape. this fact was used by phillips et al. (1999) to calibrate the dependence of the bmax–vmax and vmax– imax colours on the light curve parameter ∆m15 (b) which can, in turn, be used to separately evaluate the host galaxy extinction. using these methods for eliminating the effect of the reddening, they reanalyzed the functional form of the decline rate versus luminosity relationship and gave a value of the hubble constant of h0 = 63.3 ± 2.2 ± 3.3 km s−1 mpc−1. the use of type–ia sne is also fundamental for determining some cosmological constraints, such as ωm and ωλ that fit a λcdm models with values of 0.211±0.034 (stat) ±0.069 (sys) using a set of 252 high– redshift sne (guy et al., 2010) and 0.713+0.027−0.029 (stat) +0.036 −0.039 (sys) using a set of low–redshift nearby–hubble– flow sne (kowalski et al., 2008), respectively. in order to explore the difficult topic of the expansion of the universe it is necessary to know the evolution of metallicity in old universe that changes the hubble diagram shape. the proposed space observatory super nova acceleration probe (snap) is designed to measure the expansion of the universe and to determine the nature of the mysterious dark energy that is accelerating this expansion. snap is being proposed as part of the joint dark energy mission (jdem) (stril et al., 2010), which is a cooperative venture between nasa and the u.s. department of energy. if selected it will be launched before 2020. snap cannot achieve its main goal without progenitor/evolution solution. for comments and prospects about type–ia sn science in the decade 2010–2020 see the paper by howell et al. (2009). 6 some open questions several fundamental questions concerning cvs still remain waiting for a proper answer. we will present briefly only some of them here. one of them is the lack of a coherent classification, especially for nls. on the other hand, in gross features and in most respects, dn and nls, as well as quiescent novae, are almost indistinguishable, although, in addition to their different outbursts’ behaviour, there appear to be some further minor differences which are not yet understood (see hack & la dous 1993). the question arises of whether the outburst behaviour, the current basis of almost all classification is really a suitable criterion for sorting cvs in physically related groups. there are also too many exceptions, either systems that do not fit in any particular group or that can be included in several of them, to be able to render the observational behaviour, at least as it is used at the present, suitable. could cvs be considered simply gravimagnetic rotators? this should be the most suitable approach for studying them from a physical point of view. studies of rotational equilibria of mcvs predict that ipcvs will evolve either into pcvs or into low field strength polars — presumably unobservable, and possibly euv emitters — depending on their magnetic moments and orbital periods. indeed, there are systems, like ex hya-type, having magnetic moment similar to ipcvs above the ’period gap’ and comparable to the weakest field am her-like systems. moreover, the detection of several sw sex systems having orbital periods inside the so-called ’period gap’ opens a new interesting problem about the continuity in the evolution of cvs. the rare am canum venaticorum (am cvn) stars have extremely short orbital periods, between 10 and 65 minutes, and their spectra show no evidence for hydrogen. they appear to be helium-rich versions of cvs. they are still waiting for a general model. they are probably binary systems of two white dwarfs, but even this is still controversial. despite all the work developed during the last decades, the problem of modeling accretion discs in cvs is by no means closed, especially in quiescence. closely related is the problem of the cause of outbursts. we really do not know which of the present two families of models (disc instability models or secondary instability models) is responsible for the cvs outburst phenomenon, or in which system is each model valid, although martinez-pais et al. (1996) gave a contribution in solving this problem at least in the case of ss cygni; they found some evidence for an increase of the mass transfer rate from the secondary star as the mechanism responsible for symmetric outbursts. something similar can be said about the super-outburst phenomenon in su uma systems. gaudenzi et al. (1990), analyzing iue spectra of ss cygni, discussed about the outburst production as due to the destruction of the accretion disc. the matter, passing through the boundary layer, slowly accretes onto the wd. long and short outbursts correspond to 14 the golden age of cataclysmic variables and related objects (old and news) total or partial destruction of the disc, respectively. alternatively, could nuclear burning be responsible of the production of outbursts in cvs? indeed, nuclear burning onto white dwarf’ surface was proposed by mitrofanov (1978, 1980) as a mechanism suitable to generate x-rays in cvs. in spite of this shrewd suggestion, the community of theoreticians did not consider such a mechanism — certainly possible — worthy of taking up a part of their time. however, we believe that this alternative solution in explaining the generation of outbursts in cvs would deserve theoretician community’s care. for instance, the white dwarf surface interested in the accretion in the system ss cygni has been evaluated as 24% of the total (gaudenzi et al., 2002). there, nuclear burning could occur. accretional heating by periodic dn events increases substantially the surface temperature of the wd in cvs (godon & sion, 2002). then, the envelope thermal structure resulting from compression and irradiation should be a crucial component in understanding the envelope structure of a pre–nova wd. another problem still open is connected with the classification of cvs in three kinds, namely nmcvs, pcvs and ipcvs. this is, in our opinion, another convenient classification, although artificial, probably not necessary if cvs are studied as gravimagnetic rotators. in this way a smooth evolution of the systems could be responsible of the variations of the gravimagnetic parameters. are the ipcvs and pcvs smoothly connected via the sw sex-like systems placed just in between? sw sex systems have indeed orbital periods belong to the so-called ’period gap’, and then their presence there sure cancel that gap. could some systems behave in different ways depending on their instantaneous physical conditions? for this reason they could apparently behave sometimes as pcvs and sometimes as npcvs. an example very clear is that of ss cygni, usually classified as a non-magnetic dwarf nova. it has been detected by the integral observatory in a region of the spectrum (up to ∼ 100 kev). this emission is very hard to be explained without the presence of polar caps in the wd of the system. several proofs have been shown and discussed many times by giovannelli’s group in order to demonstrate the intermediate polar nature of it (e.g., giovannelli, 1996, and references therein; giovannelli & sabau-graziati, 1998; 2012a); indeed, ss cygni shows characteristics of a nmcv, as well as those of ip and sometimes even those of polars, although its position in the log pspin–log porb plane is very close to the line where ips lie. important results are coming from the spitzer space telescope with the detection of an excess (3-8) µm emission from magnetic cvs, due to dust (howell et al., 2006; brinkworth et al., 2007). gaudenzi et al. (2011) discussed about the reasons of the variable reddening in ss cyg and demonstrated that this reddening is formed by two components: the first is interstellar in origin, and the second (intrinsic to the system itself) is variable and changes during the evolution of a quiescent phase. moreover, an orbital modulation also exists. the physical and chemical parameters of the system are consistent with the possibility of formation of fullerenes. the spitzer space telescope detected the presence of fullerenes in a young planetary nebula (cami et al., 2010). fullerenes are the first bricks for the emergence of the life. therefore, the possible presence of fullerenes in cvs opens a new line of investigation, foreboding of new interesting surprises. 7 conclusions at the end of this review it appears evident that the most suitable approach for studying cvs from a physical point of view is to consider them as gravimagnetic rotators. the detection of several sw sex systems having orbital periods inside the so-called ’period gap’ opens a new interesting problem about the continuity in the evolution of cvs. are the ipcvs and pcvs smoothly connected via the sw sex-like systems placed just in between? in order to fully understand the emission properties and evolution of cvs, the mass–transfer process needs to be clearly understood, especially magnetic mass transfer, as well as the properties of magnetic viscosity in the accretion discs around compact objects. consequently, the investigation on the magnetic field intensities in wds appears crucial in understanding the evolution of cvs systems, by which it is possible to generate classical novae (e.g., isern et al., 1997) and type-ia supernovae (e.g., isern et al., 1993). in those catastrophic processes the production of light and heavy elements, and then the knowledge of their abundances provides strong direct inputs for cosmological models and cosmic ray generation problems. acknowledgments this research has made use of nasa’s astrophysics data system. references [1] andronov, i.l., chinarova, l.l., han, w., kim, y., yoon, j.-n.: 2008, a&a, 486, 855. 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[172] zhang, e., robinson, e.l., stiening, r.f., horne, k.: 1995, apj, 454, 447. 20 http://dx.doi.org/10.1038/nature06602 http://dx.doi.org/10.1093/mnras/227.1.23 http://dx.doi.org/10.1017/cbo9780511586491 http://dx.doi.org/10.1007/bf00613240 http://dx.doi.org/10.1007/bf00645229 http://dx.doi.org/10.1063/1.1518170 http://dx.doi.org/10.1086/381742 introduction multifrequency emissions renewed interest for cataclysmic variables classical and recurrent novae progenitors of sn ia some open questions conclusions 1 acta polytechnica ctu proceedings 1(1): 1–12, 2014 1 doi: 10.14311/app.2014.01.0001 old and new from multifrequency astrophysics franco giovannelli1, lola sabau-graziati2 1inaf istituto di astrofisica e planetologia spaziali, area di ricerca di tor vergata, via fosso del cavaliere, 100 i00177 roma, italy 2intadpt. cargas utiles y ciencias del espacio, c/ra de ajalvir, km 4 e28850 torrejón de ardoz, madrid, spain corresponding author: franco.giovannelli@iaps.inaf.it abstract in this short review paper we comment on some the most important steps that have been made in the past decades for a better understanding of the physics governing our universe. the results we discuss come from the many groundand-space-based experiments developed for measuring astrophysical sources in various energy bands. these experimental results are discussed within the framework of current theoretical models. because of the limited length of this paper, we have selected only a few topics that, in our opinion, have been crucial for the progress of our understanding of the physics of cosmic sources. keywords: multifrequency astrophysics. 1 introduction with the advent of space-based experiments, it has been demonstrated that cosmic sources emit energy practically across the entire electromagnetic spectrum, albeit from different physical processes. several observations stand as witness to these processes. since the observed fluxes from cosmic sources can be highly variable in time and frequency, it follows that the physical processes from which they originate are in themselves highly variable. therefore simultaneous multifrequency observations are strictly necessary in order to understand the actual behaviour of cosmic sources. indeed, space experiments have opened practically all of the electromagnetic “windows” on the universe. a discussion of the most important results coming from multifrequency photonic astrophysics experiments will provide new inputs for the advance of our knowledge of physics, very often in the most extreme physical conditions. we remark on the sheer magnitude of the high quality data across practically the whole electromagnetic spectrum that has become available to the scientific community since the beginning of the space era. with these data, we are attempting to explain the physics governing the universe, and, moreover, its origin, which has been and still is a matter of the greatest curiosity for humanity. we know for certain that the universe has an absolute power limit lmax ∼ �pl/tpl ∼ c5/g ∼ 3.6×1059 erg s−1, where �pl and tpl are the planck energy and planck time, respectively, c the light velocity and g the gravitational constant. this amount of power is produced in different kind of cosmic sources, namely, the early universe (eu), quasars (qsos) and active galactic nuclei (agns), supernovae (sne), neutron stars (nss), and black holes (bhs), all types of galaxies with their stars and interstellar medium (ism), and intergalactic medium (igm) (e.g., lipunov, 1995). they radiate particles and photons at different levels of energy across the entire electromagnetic spectrum from their origins. however, the cosmic particle radiation arriving near earth, with energies from ∼ 106 to ∼ 1020 ev, is nearly isotropic, because of the galactic magnetic field, which cancels any particular directionality in the galaxy. such cosmic particle radiation apparently includes the nuclei of all known elements, as well as electrons, positrons, and antiprotons. so in spite of the cosmic rays being carriers of rich astrophysical information, it is very difficult to understand their message clearly. indeed, although it is evident that they must originate in different sources, it is at the same time extremely difficult to separate the different contributions. for more complete discussions, the reader can note the substantive books on this topic, containing excellent reviews, published by the italian physical society (giovannelli & mannocchi, 1989, 1991, 1993, 1995, 1997, 1999, 2001, 2003, 2005, 2007, 2009, 2011), by the italian astronomical society (giovannelli & sabau-graziati, 1996; 1999, 2002a, 2002b, 2010a, 2012a), by the chinese journal of astronomy and astrophysics (giovannelli & sabau-graziati, 2003, 2006, 2008), and by acta poly1 http://dx.doi.org/10.14311/app.2014.01.0001 franco giovannelli, lola sabau-graziati technica (giovannelli & mannocchi, 2013; giovannelli & sabau-graziati, 2014). with “passive” physics experiments (.i.e., observations) we view our universe, while with active physics experiments we try to reproduce some of the physical conditions and processes occurring somewhere in the universe. both kinds of experiments converge to the knowledge of the physics governing the universe. in this paper we will discuss what seem to us some of the most relevant results obtained in the recent past that significantly improve our knowledge of the physics governing our universe. deeper discussions about astroparticle physics can be found in the review papers by giovannelli (2007, 2009, 2011, 2013). giovannelli & sabau-graziati (2010b, 2012b), and de angelis, mansutti & persic (2008) discussed in their review papers the multifrequency behaviour of high energy cosmic sources, and very high energy (vhe) γ-ray astrophysicsal sources. 1.1 astroparticle physics development the subject of high energy astrophysics is generally approached through the study of cosmic rays. the reason for this is historical in nature. since the discovery of this extraterrestrial radiation by victor hess (1912), the scientific research involved in trying to discover the nature nature of these sources has been extensive. as a result, many separate research fields have been developed. before particle accelerators came into operation, high energy cosmic rays were the laboratory tools for investigations of elementary particle production, and to date they are still the only source of particles with energies greater than 1012 ev. the research into the composition of the radiation led to the developing study of the astrophysical environment using the information in the charge, mass, and energy spectra; this field is also known as particle astrophysics. now, the large hadronic collider (lhc), described by straessner et al. (2011), is able to reach tev energies for p-p interactions, and has attained energyies of 7 tev, in order to search for the higgs’ boson with the atlas detector (aad et al., 2012). no significant excess of events over the expected background is observed and limits on the higgs boson production cross section are derived for a higgs boson mass in the range 110 gev < mh < 300 gev. the observations exclude the presence of a standard model higgs boson with a mass 145 < mh < 206 gev at 95% confidence level. of great importance was the discovery of high energy photons near the top of the earth’s atmosphere. this originated the development of new astronomical fields such as the x-ray or γ-ray astronomy. but many of these high energy photons have their origin in the interactions of the high energy charged particles with cosmic matter, light, or magnetic fields. the particle astrophysics and the astronomical research fields have found in this fact a bond to join their efforts in trying to understand the high energy processes which occur in astrophysical systems. a summary on the status of the search for the origin of the highest energy cosmic rays has been published by biermann (1999). he mentioned several competing proposals, such as the supersymmetric particles. biermann notes that gamma ray bursts must also give rise to energetic protons, interacting high energy neutrinos and cosmological defects. in his paper, biermann discussd the propagation of these particles, assuming that they are charged, and concluding that the distribution of arrival directions of the highest energy particles on the sky ought to reflect the source distribution as well as the propagation history. he remarked that the present status of our observations can be summarized as inconclusive. however, he concluded as follows: if we can identify the origin of the events at the highest energies, beyond 5×1019 ev, the greisen–zatsepin–kuzmin cutoff due to the microwave background, near to 1021 ev, and if we can establish the nature of their propagation through the universe to us, then we will obtain a tool to do physics at eev energies. the arrival directions of ≥ 60 eev ultra-high-energy cosmic rays (uhecrs) cluster along the supergalactic plane and correlate with active galactic nuclei (agn) within ≈ 100 mpc (abraham et al., 2007, 2008). the association of several events with the nearby radio galaxy cen a supports the paradigm that uhecrs are powered by supermassive black-hole engines and accelerated to ultra-high energies in the shocks formed by variable plasma winds in the inner jets of radio galaxies. the gzk horizon length of 75 eev uhecr protons is ≈ 100 mpc, so that the auger results are consistent with an assumed proton composition of the uhecrs. in this scenario, the sources of uhecrs are fr ii radio galaxies and fr i galaxies like cen a with scattered radiation fields that enhance uhecr neutral–beam production. radio galaxies with jets pointed away from us (i.e., more toward the plane of the sky) can still be observed as uhecr sources due to deflection of uhecrs by magnetic fields in the radio lobes of these galaxies. a broadband ∼ 1 mev–10 eev radiation component in the spectra of blazar agn is formed by uhecr–induced cascade radiation in the extragalactic background light. this emission is too faint to be seen from cen a, but could be detected from more luminous blazars (dermer et al., 2009). recent evidence from the pierre auger observatory suggests a transition, at 5 eev-10eev in the composition of ultra high energy cosmic rays (uhecrs), from protons to heavier nuclei such as iron (abraham et al., 2010). piran (2010) considered the implications of 2 old and new from multifrequency astrophysics the heavier composition on the sources of uhecrs. he concluded that with typical reasonable parameters of a few ng for the extra–galactic magnetic field (egmf) intensity and a coherence distance of a mpc the distance that nuclei uhecr above the gzk energy traverses before photodisintegrating is only a few mpc. in spite of the significantly weaker limits on the luminosity, cen a is the only currently active potential source of nuclei uhecrs within this distance. the large deflections erases the directional anisotropy expected from a single source. if indeed the composition of above gzk– uhecrs is iron and if the egmf is not too small then cen a is the dominant source of observed nuclei uhecrs above the gzk limit. in summary, charged cosmic rays are influenced in their propagation through space by the magnetic fields in the galaxy, and for the lowest energy particles also in the solar system. the result is that the distribution of arrival directions as the radiation enters the earth’s atmosphere is nearly isotropic. it is not possible to identify the sources of the cosmic rays by detecting them. however, in the high energy interactions produced at the source, electrically neutral particles such as photons, neutrons, and neutrinos are also produced and their trajectories are not deviated, being directed from their point of origin to the observer. owing to their short lifetime, neutrons cannot survive the path length to the earth (decay length ∼ 9 pc at 1 pev) and neutrinos do not interact efficiently in the atmosphere. it is in this context that the gamma ray astronomy has demonstrated itself to be a powerful tool. the observations made to date have detected γ-rays from many astronomical objects such as neutron stars, interstellar clouds, the center of our galaxy and the nuclei of active galaxies (agns). one might expect very important implications for high energy astrophysics from the observations at energies greater than 1011 ev of extragalactic sources (e.g., hillas & johnson, 1990). the fluxes of γ-rays at these energies are attenuated because of their interactions with the cosmic radio, microwave, infrared and optical radiation fields. measurements of the flux attenuation can then provide important information on the distribution of such fields. for instance, the threshold energy for pair production in reactions of photons with the 2.7 k background radiation is reached at 1014 ev and the absorption length is of the order of ∼ 7 kpc. for the infrared background the maximum absorption is reached at energies greater than 1012 ev. the qualitative problem of the origin of cosmic rays is practically solved, while the quantitative problem in determining the fraction of them coming from the different possible sources is still open. 2 very high energy sources the most exciting results of the last decade have been obtained in the field of vhe astrophysics from different experiments (e.g. cgro/egret, wipple, hegra, cangaroo, celeste, stacee, tibet, hess, veritas, milagro, magic) that detected many vhe cosmic sources. the high energy sky. with the exception of crab nebula, vela x, and 3c 273, was empty until middle nineties. “fast forward” to 19th april 2012, the vhe sky (e > 100 gev) is populated by 107 cosmic sources: 46 out of 107 extragalactic and 61 galactic (http://www.mppmu.mpg.de/∼rwagner/sources/ or http://tevcat.uchicago.edu). one of the most interesting results has been the determination of the spectral energy distribution (sed) of the crab nebula, thanks to many measurements obtained by different he–vhe experiments (albert et al., 2008b). another exciting result has been the detection of the first variable galactic tev source, namely the binary pulsar psr b1259-63 (aharonian et al., 2005). they found that the radio silence occurs during the time in which the pulsar is occulted by the excretion disk of the be star. the many detected sources, representing different galactic and extragalactic source populations, are supernova remnants (snrs), pulsar wind nebulae (pwne), giant molecular clouds (gmcs), star formation regions (sfrs), compact binary systems (cbss), and active galactic nuclei (agns). paredes & persic (2010) reviewed the results from magic cherenkov telescope for most of the former class of sources. models of tev agns have been discussed by lenain (2010). 3 diffuse extragalactic background radiation after the big bang, the universe started to expand with a fast cooling. the cosmic radiation observed now is probably an admixture of different components which had their origin in different stages of the evolution as the results of different processes. this is the diffuse extragalactic background radiation (debra), which, if observed in different energy ranges, allows the study of many astrophysical, cosmological, and particle physics phenomena (ressell & turner, 1990. debra is the witness of the whole history of the universe from the big bang to present time. such history is marked by three main experimental witnesses supporting the big bang theory (e.g. giovannelli & sabau-graziati, 2008): the light element abundances (burles, nollett & turner, 2001); the cmbr temperature at various redshifts as determined by srianand, petitjean & ledoux (2000), and 3 franco giovannelli, lola sabau-graziati the references therein; the cmb at z = 0 as result of cobe (tcmbr(0) = 2.726 ± 0.010 k), which is well fitted by a black body spectrum (mather et al., 1994). at z' 2.34, the cmbr temperature is: 6.0 k < tcmbr(2.34) < 14.0 k. the prediction from the hot big bang: tcmbr = tcmbr(0) × (1 + z) gives tcmbr(2.34) = 9.1 k, which is consistent with the measurement (srianand, petitjean & ledoux, 2000). 4 reionization of the universe after the epoch of recombination (last scattering) between ≈ 3.8 × 105− ≈ 2 × 108 yr (z ≈ 1000 − 20), the universe experienced the so–called dark ages, where the dark matter halos collapsed and merged until the appearance of the first sources of light. this ended the dark ages. the ultraviolet light from the first sources of light also changed the physical state of the gas (hydrogen and helium) that fills the universe from a neutral state to a nearly fully ionized one. this was the reionization era where the population iii stars formed, and consequently, the first sne and grbs. this occurred between ≈ (2 − 5) × 108 yr (z ≈ 20 − 10). soon after population ii stars started to form and probably the second wave of reionization occurred and stopped at ≈ 9 × 108 yr (z ≈ 6) after the big bang, and then the evolution of galaxies started (e.g. djorgovski, 2004, 2005). quasars – the brightest and most distant objects known – offer a window on the reionization era, because neutral hydrogen gas absorbs their ultraviolet light. reionization drastically changes the environment for galaxy formation and evolution and in a hierarchical clustering scenario, the galaxies responsible for reionization may be the seeds of the most massive galaxies in the local universe. reionization is the last global phase transition in the universe. the reionization era is thus a cosmological milestone, marking the appearance of the first stars, galaxies and quasars. recent results obtained by ouchi et al. (2010) give an important contribution for solving such a problem. indeed, from the the lyα luminosity function (lf), clustering measurements, and lyα line profiles based on the largest sample to date of 207 lyα emitters at z = 6.6 on the 1 deg2 sky of subaru/xmm-newton deep survey field, ouchi et al. (2010) found that the combination of various reionization models and observational results about the lf, clustering, and line profile indicates that there would exist a small decrease of the intergalactic medium’s (igm’s) lyα transmission owing to reionization, but that the hydrogen igm is not highly neutral at z = 6.6. their neutral-hydrogen fraction constraint implies that the major reionization process took place at z >∼ 7. the w. m. keck 10-m telescope has shown the quasar sdss j1148+5251 at a redshift of 6.41 (≈ 12.6× 109 yr ago). this is currently the most distant quasar known (djorkovski, 2004). this measurement does not contradict the result found for the epoch of reionization. however, the search of the epoch of reionization is still one of the most important open problems for understanding the formation of the first stars, galaxies and quasars. 5 clusters of galaxies the problems of the production and transport of heavy elements seems to have been resolved. indeed, thermally driven galactic winds, such as from m82, have shown that only active galaxies with an ongoing starburst can enrich the icm with metals. the amounts of metals in the icm is at least as high as the sum of the metals in all galaxies of the cluster (e.g. tozzi et al., 2003). several clusters of galaxies, having strong radio emission, have been associated with egret sources. this is an important step in clarifying the nature of many unknown egret sources (colafrancesco, 2002). however, in the first 11 months of operations of the fermi lat monitoring program of cgs no γ-ray emission from any of the monitored cgs has been detected (ackermann et al., 2010b). in spite of many important results coming from satellites of the last decade, the hierarchical distribution of the dark matter, and the role of the intergalactic magnetic fields in cgs are still open. simultaneous multifrequency measurements with higher sensitivity instruments, in particular those in hard x-ray and radio energy regions and optical-to-near infrared (nir) could solve such problems. 6 dark energy and dark matter by using different methods to determine the mass of galaxies it has been found a discrepancy that suggests ∼ 95% of the universe is in a form that cannot be seen. this form of unknown content of the universe is the sum of dark energy (de) and dark matter (dm). colafrancesco (2003) deeply discussed about new cosmology. the discovery of the nature of the dark energy may provide an invaluable clue for understanding the nature and the dynamics of our universe. however, there is ∼ 30% of the matter content of the universe which is dark and still requires a detailed explanation. baryonic dm consisting of machos (massive astrophysical compact halo objects) can yield only some fraction of the total amount of dark matter required by cmb observations. wimps (weakly interacting massive particles) (non-baryonic dm) can yield the needed cosmological amount of dm and its large scale distribution provided that it is “cold” enough. several options have been proposed so far like: i) light neutrinos with mass in 4 old and new from multifrequency astrophysics the range mν ∼ 10−30 ev, ii) light exotic particles like axions with mass in the range maxion ∼ 10−5−10−2 ev or weakly interacting massive particles like neutralinos with mass in the range mχ ∼ 10 − 1000 gev, this last option being favored at present (see, e.g., ellis 2002). eros and macho, two experiments based on the gravitational microlensing, were developed. two lines of sight have been probed intensively: the large (lmc) and the small (smc) magellanic clouds, located 52 kpc and 63 kpc respectively from the sun (palanque-delabrouille, 2003). with 6 years of data towards the lmc, the macho experiment published a most probable halo fraction between 8 and 50% in the form of 0.2 m� objects (alcock et al., 2000). most of this range is excluded by the eros exclusion limit, and in particular the macho preferred value of 20% of the halo. among experiments for searching wimps as dark matter candidates, there is pamela, an experiment devoted to a search for dark matter annihilation, antihelium (primordial antimatter), new matter in the universe (strangelets?), the study of cosmic-ray propagation (light nuclei and isotopes), electron spectrum (local sources?), solar physics and solar modulation, and terrestrial magnetosphere. a comparison of the expected pamela results with many other experiments has been discussed by morselli (2007). bruno (2011) discussed some results from pamela. the search for dm is one of the main open problems of today’s astroparticle physics. 7 the galactic center the galactic center (gc) is one of the most interesting places for testing theories in which frontier physics plays a fundamental role. there is an excellent review of mezger, duschl & zylka (1996), which discusses the physical state of stars and interstellar matter in the galactic bulge (r ∼ 0.3–3 kpc from the dynamic center of the galaxy), in the nuclear bulge (r < 0.3 kpc) and in the sgr a radio and gmc complex (the central ∼ 50 pc of the milky way). this review reports also a list of review papers and conference proceedings related to the galactic center with bibliographic details. in the review paper by giovannelli & sabau-graziati (2004, and the references therein) the multifrequency behaviour of the galactic center has been also discussed. larosa et al. (2000) presented a wide-field, high dynamic range, high-resolution, long-wavelength (λ = 90 cm) vla image of the galactic center region. this is the most accurate image of the gc. while highly obscured in optical and soft x-rays; it shows a central compact object (a black hole candidate) with m ∼ 3.6 × 106 m� (genzel et al., 2003a), which coincides with the compact radio source sgr a∗ [r.a. 17 45 41.3 (hh mm ss); dec.: -29 00 22 (dd mm ss)]. sgr a∗ in x-rays/infrared is highly variable (genzel et al., 2003b). the gc is also a good candidate for indirect dark matter observations. moreover, the detected excess of he γ-rays at gc would be produced by neutralino annihilation in the dark matter halo. such an excess could be better measured by the fermi observatory. 8 gamma-ray bursts the many theoretical descriptions of gamma-ray bursts (grbs) show that the origin of these sources is still an open and strongly controversial topic. fireball (fb) model (meszaros & rees, 1992; piran, 1999), cannon ball (cb) model (dar & de rújula, 2004), spinninprecessing jet (spj) model (fargion, 2003a,b; fargion & grossi, 2006), fireshell (izzo et al., 2010) model — directly coming from electromagnetic black hole (embh) model (e.g. ruffini et al. 2003 and the references therein) — are the most popular, but each one against the others. important implications on the origin of the highest redshift grbs are coming from the detection of the grb 080913 at z =6.7 (greiner et al., 2009), grb 090423 at z ∼ 8.2 (tanvir et al., 2009), and grb 090429b (cucchiara et al., 2011). this means that really we are approaching to the possibility of detecting grbs at the end of dark era, where the first pop iii stars appeared. izzo et al. (2010) discussed successfully a theoretical interpretation of the grb 090423 within their fireshell model. wang & dai (2009) studied the high-redshift star formation rate (sfr) up to z ' 8.3 considering the swift grbs tracing the star formation history and the cosmic metallicity evolution in different background cosmological models including λcdm, quintessence, quintessence with a time-varying equation of state and brane-world models. λcdm model is the preferred which is however compared with other results. although great progress has been obtained in the last few years, grbs theory needs further investigation in the light of the experimental data coming from old and new satellites, often coordinated, such as bepposax or batse/rxte or asm/rxte or ipn or hete or integral or swift or agile or fermi or maxi. 9 extragalactic background light space is filled with diffuse extragalactic background light (ebl) which is the sum of starlight emitted by galaxies through the history of the universe. high energy γ-rays traversing cosmological distances are expected to be absorbed through their interactions with the ebl by: γvhe + γebl −→ e+ e−. then the γ-ray flux φ is suppressed while travelling from the emission 5 franco giovannelli, lola sabau-graziati point to the detection point, as φ = φ0e −τ(e,z), where τ(e,z) is the opacity. the e–fold reduction [τ(e,z) = 1] is the gamma ray horizon (grh) (e.g. blanch & martinez, 2005; martinez, 2007). the direct measurement of the ebl is difficult at optical to infrared wavelengths because of the strong foreground radiation originating in the solar system. however, the measurement of the ebl is important for vhe gamma-ray astronomy, as well as for astronomers modelling star formation and galaxy evolution. second only in intensity to the cosmic microwave background (cmb), the optical and infrared (ir) ebl contains the imprint of galaxy evolution since the big bang. this includes the light produced during formation and reprocessing of stars. current measurements of the ebl are reported in the paper by schroedter (2005, and references therein). he used the available vhe spectra from six blazars. more recently, the redshift region over which the gamma reaction history (grh) can be constrained by observations has been extended up to z = 0.536. upper ebl limit based on 3c 279 data have been obtained (albert et al., 2008a). the universe is more transparent to vhe gamma rays than expected. thus many more agns could be seen at these energies. indeed, abdo et al. (2009a) observed a number of tev-selected agns during the first 5.5 months of observations with the large area telescope (lat) on–board the fermi gamma-ray space telescope. redshift– dependent evolution is detected in the spectra of objects detected at gev and tev energies. the most reasonable explanation for this is absorption on the ebl, and as such, it would represent the first model–independent evidence for absorption of γ-rays on the ebl. abdo et al. (2010b) by using a sample of γ-ray blazars with redshift up to z ∼ 3, and grbs with redshift up to z ∼ 4.3, measured by fermi/lat placed upper limits on the γ-ray opacity of the universe at various energies and redshifts and compared this with predictions from well–known ebl models. they found that an ebl intensity in the optical-ultraviolet wavelengths as great as predicted by the ”baseline” model of stecker, malkan & scully (2006) that can be ruled out with high confidence. 10 relativistic jets relativistic jets have been found in numerous galactic and extragalactic cosmic sources at different energy bands. the emitted spectra of jets from cosmic sources of different nature are strongly dependent on the angle formed by the beam axis and the line of sight, and obviously by the lorentz factor of the particles (e.g. bednarek et al., 1990 and the references therein; beall, guillory & rose, 1999, 2009; beall, 2002, 2003, 2008, 2009; beall et al., 2006, 2007). so, observations of jet sources at different frequencies can provide new inputs for the comprehension of such extremely efficient carriers of energy, like for the cosmological grbs. the discovered analogy among µ–qsos, qsos, and grbs is fundamental for studying the common physics governing these different classes of objects via µ–qsos, which are galactic, and then apparently brighter and with all processes occurring in time scales accessible by our experiments (e.g. chaty, 1998). chaty (2007) remarked the importance of multifrequency observations of jet sources by means of the measurements of grs 1915+105. dermer et al. (2009) suggest that ultra-high energy cosmic rays (uhecrs) could come from black hole jets of radio galaxies. spectral signatures associated with uhecr hadron acceleration in studies of radio galaxies and blazars with fermi observatory and ground–based γ-ray observatories can provide evidence for cosmic-ray particle acceleration in black hole plasma jets. also in this case, γ-ray multifrequency observations (mev–gev–tev) together with observations of pev neutrinos could confirm whether black-hole jets in radio galaxies accelerate the uhecrs. despite their frequent outburst activity, microquasars have never been unambiguously detected emitting high-energy gamma rays. the fermi/lat has detected a variable high-energy source coinciding with the position of the x-ray binary and microquasar cygnus x-3. its identification with cygnus x-3 is secured by the detection of its orbital period in gamma rays, as well as the correlation of the lat flux with radio emission from the relativistic jets of cygnus x-3. the γ-ray emission probably originates from within the binary system (abdo et al., 2009b). also the microquasar ls 5039 has been unambiguously detected by fermi/lat being its emission modulated with a period of 3.9 days. analyzing the spectrum, variable with the orbital phase, and having a cutoff, abdo et al. (2009c) conclude that the γ-ray emission of ls 5039 is magnetospheric in origin, like that of pulsars detected by fermi. this experimental evidence of emission in the gev region from microquasars opens an interesting window about the formation of relativistic jets. 11 cataclysmic variables the detection of cvs with the integral observatory (barlow et al., 2006) have recently renewed the interest of high energy astrophysicists for such systems, and subsequently involving once more the low–energy astrophysical community. the detection of cvs having orbital periods inside the so-called period gap between 2 and 3 hours, which separates polars (apparently generating gravitational radiation) from intermediate polars (which suffer magnetic braking) renders attractive 6 old and new from multifrequency astrophysics the idea of the physical continuity between these two classes. further investigations are necessary for solving this important problem. for a recent review on cvs see the paper by giovannelli & sabau-graziati (2012c). 12 high mass x-ray binaries for general reviews see e.g. giovannelli & sabaugraziati (2001, 2004) and van den heuvel (2009) and references therein. hmxbs are young systems, with age ≤ 107 yr, mainly located in the galactic plane (e.g., van paradijs, 1998). a compact object — the secondary star —, mostly a magnetized neutron star (x-ray pulsar) is orbiting around an early type star (o, b, be) — the primary — with m ≥ 10 m�. the optical luminosity of the system is dominated by the early type star. such systems are the best laboratory for the study of accreting processes, thanks to their relatively high luminosity in a large part of the electromagnetic spectrum. because of the strong interactions between the optical companion and collapsed object, low and high energy processes are strictly related. in x-ray/be binaries the mass loss processes are due to the rapid rotation of the be star, the stellar wind and, sporadically, to the expulsion of casual quantity of matter essentially triggered by gravitational effects close to the periastron passage of the neutron star. the long orbital period (> 10 days) and a large eccentricity of the orbit (> 0.2) together with transient hard x-ray behavior are the main characteristics of these systems. among the whole sample of galactic systems containing 114 x-ray pulsars (johnstone, 2005), only few of them have been extensively studied. among these, the system a 0535+26/hde 245770 is the best known thanks to concomitant favorable causes, which rendered possible thirty eight years of coordinated multifrequency observations, most of them discussed by e.g. giovannelli & sabau-graziati (1992, 2008), burger et al. (1996). accretion powered x-ray pulsars usually capture material from the optical companion via stellar wind, since this primary star generally does not fill its roche lobe. however, in some specific conditions (e.g. the passage at the periastron of the neutron star) and in particular systems (e.g. a 0535+26/hde 245770), it is possible the formation of a temporary accretion disk around the neutron star behind the shock front of the stellar wind. this enhances the efficiency of the process of mass transfer from the primary star onto the secondary collapsed star, as discussed by giovannelli & ziolkowski (1990) and by giovannelli et al. (2007) in the case of a 0535+26. giovannelli & sabau-graziati (2011) discussed the history of the discovery of optical indicators of high energy emission in the prototype system a0535+26/hde 245770 ≡ flavia’ star, updated to the march–april 2010 event when a strong optical activity occurred roughly 8 days before the x-ray outburst (caballero et al., 2010) that was predicted by giovannelli, gualandi & sabaugraziati (2010). this event together with others occurred in the past allowed to giovannelli & sabaugraziati (2011) to conclude that x-ray outbursts occur ∼ 8 days after the periastron passage. giovannelli, bisnovatyi-kogan & klepnev (2013) developed a model for explaining such a delay by the time of radial motion of the matter in a non–stationary accretion disk around the neutron star, after an increase of the mass flux in the vicinity of a periastral point in the binary. this time is determined by the turbulent viscosity, with the parameter α = 0.1 − 0.3. however how x-ray outbursts are triggered in xray pulsars constitute one important still open problem giving rise to controversy within astrophysicists. important news are coming also from gev observations of hmxbs. indeed, abdo et al. (2009e) present the first results from the observations of lsi + 61◦303 using fermi/lat data obtained between 2008 august and 2009 march. their results indicate variability that is consistent with the binary period, with the emission being modulated at 26.6 days. this constitutes the first detection of orbital periodicity in high–energy γ-rays (20 mev-100 gev). the light curve is characterized by a broad peak after periastron, as well as a smaller peak just before apastron. the spectrum is best represented by a power law with an exponential cutoff, yielding an overall flux above 100 mev of ' 0.82×10−6 ph cm−2 s−1, with a cutoff at ∼ 6.3 gev and photon index γ ∼ 2.21. there is no significant spectral change with orbital phase. the phase of maximum emission, close to periastron, hints at inverse compton scattering as the main radiation mechanism. however, previous very high-energy gamma ray (> 100 gev) obser vations by magic and veritas show peak emission close to apastron. this and the energy cutoff seen with fermi suggest that the link between he and vhe gamma rays is nontrivial. this is one open problem to be solved in future. 12.1 obscured sources and supergiant fast x-ray transients relevant are integral results about a new population of obscured sources and supergiant fast x-ray transients (sfxts) (chaty & filliatre, 2005; chaty, 2007; rahoui et al., 2008; chaty, 2008). the importance of the discovery of this new population is based on the constraints on the formation and evolution of hmxbs: does dominant population of short-living systems – born with two very massive components – oc7 franco giovannelli, lola sabau-graziati cur in rich star-forming region? what will happen when the supergiant star dies? are primary progenitors of ns/ns or ns/bh mergers good candidates of gravitational waves emitters? can we find a link with short/hard γ-ray bursts? 13 ultra–compact double–degenerated binaries ultra-compact, double-degenerated binaries (ucd) consist of two compact stars, which can be black holes, neutron stars or white dwarfs. in the case of two white dwarfs revolving around each other with an orbital period porb ≤ 20 min. the separation of the two components for a ucd with porb ≈ 10 min or shorter is smaller than jupiter’s diameter. these ucd are evolutionary remnants of low–mass binaries, and they are numerous in the milky way. the discovery of ucd is foreboding interesting hints for gravitational–wave possible detection with lisa observatory. 14 magnetars the discovery of magnetars (anomalous x-ray pulsars – axps – and soft gamma-ray repeaters – sgrs) is also one of the most exciting results of the last years (mereghetti & stella, 1995; van paradijs, taam & van den heuvel, 1995; and e.g. review by giovannelli & sabau-graziati, 2004 and the references therein). indeed, with the magnetic field intensity of order 1014 − 1015 g a question naturally arises: what kind of sn produces such axps and sgrs? are really the collapsed objects in axps and sgrs neutron stars? (e.g. hurley, 2008). with such high magnetic field intensity an almost ‘obvious’ consequence can be derived: the correspondent dimension of the source must be of ∼ 10 m (giovannelli & sabau-graziati, 2006). this could be the dimension of the acceleration zone in supercompact stars. could they be quark stars? ghosh (2009) discussed some of the recent developments in the quark star physics along with the consequences of possible hadron to quark phase transition at high density scenario of neutron stars and their implications on the astroparticle physics. important consequences could be derived by considering the continuity among rotation-powered pulsars, magnetars, and millisecond pulsars. such continuity has been experimentally demonstrated (kuiper, 2007). however, the physics underlying that observational continuity is not yet clear. 15 neutrino astronomy for a short discussion about neutrino astronomy, see for instance the paper by giovannelli (2007 and the references therein), as well as all the papers of the session neutrino astronomy, which appeared in the proceedings of the vulcano workshops 2006, 2008, and 2010 (giovannelli & mannocchi, 2007, 2009, 2011). however, it is important to remark that several papers have appeared about: i) the sources of he neutrinos (aharonian, 2007) and diffuse neutrinos in the galaxy (evoli, grasso & maccione, 2007); ii) potential neutrino signals from galactic γ-ray sources (kappes et al., 2007); iii) galactic cosmic-ray pevatrons with multi-tev γ-rays and neutrinos (gabici & aharonian, 2007); iv) results achieved with amanda: 32 galactic and extragalactic sources have been detected (xu & the icecube collaboration, 2008); diffuse neutrino flux from the inner galaxy (taylor et al., 2008); discussion about vhe neutrino astronomic experiments (cao, 2008). important news and references can be found in the proceedings of the les rencontres de physique de la vallée d’aoste (greco, 2009, 2010). news about the neutrino oscillations have been reported by mezzetto (2011). the angle θ13 is different than zero: sin2 θ13 = 0.013. this result opens the door to cp violation searches in the neutrino sector, with profound implications for our understanding of the matter–antimatter asymmetry in the universe. 16 conclusions and reflections it is becoming increasingly clear that the energy régime covered by vhe γ-ray astronomy will be able to address a number of significant scientific questions, which include: i) what parameters determine the cut-off energy for pulsed γ-rays from pulsars? ii) what is the role of shell-type supernovae in the production of cosmic rays? iii) at what energies do agn blazar spectra cut-off? iv) are gamma blazar spectral cut-offs intrinsic to the source or due to intergalactic absorption? v) is the dominant particle species in agn jets leptonic or hadronic? vi) can intergalactic absorption of the vhe emission of agn’s be a tool to calibrate the epoch of galaxy formation, the hubble parameter, and the distance to γ-ray bursts? vii) are there sources of γ-rays which are ‘loud’ at vhes, but ‘quiet’ at other wavelengths? it appears evident the importance of multifrequency astrophysics. there are many problems in performing simultaneous multifrequency, multienergy multisite, multiinstrument, multiplatform measurements due to: i) objective technological difficulties; ii) sharing common scientific objectives; iii) problems of scheduling and budgets; iv) politic management of science. in spite of the many ground-based and space-based experiments providing an impressive quantity of excellent data in different energy regions, many open problems still exist. we believe that only by drastically 8 old and new from multifrequency astrophysics changing the philosophy of the experiments will it be possible to solve most of the present open problems. for instance, in the case of space–based experiments, small satellites, dedicated to specific missions and problems, and having the possibility of scheduling very long time observations, must be supported because of their relative faster preparation, easier management and lower costs with respect to medium and large satellites. we strongly believe that in the next decades “passive” physics experiments in space, as well as groundbased, and perhaps lunar-based observatories will be the most suitable probes in sounding the physics of the universe. probably the active physics experiments have already reached the maximum dimensions compatible with a reasonable cost/benefit ratio, with the obvious exception of the neutrino astronomy experiments. acknowledgement we wish to thank the soc of the karlovy vary 10th integral/bart workshop for inviting us to discuss this review during the workshop, and the loc for logistical support for one of us (fg). this research has made use of nasa’s astrophysics data system. references [1] aad, g. et al. 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(icecube collaboration): 2008, n. phys. b 175-176, 401. 12 http://dx.doi.org/10.1063/1.1587092 http://dx.doi.org/10.1086/431173 http://dx.doi.org/10.1038/35050020 http://dx.doi.org/10.1038/nature08459 http://dx.doi.org/10.1063/1.3076687 http://dx.doi.org/10.1086/376731 introduction astroparticle physics development very high energy sources diffuse extragalactic background radiation reionization of the universe clusters of galaxies dark energy and dark matter the galactic center gamma-ray bursts extragalactic background light relativistic jets cataclysmic variables high mass x-ray binaries obscured sources and supergiant fast x-ray transients ultra–compact double–degenerated binaries magnetars neutrino astronomy conclusions and reflections 178 acta polytechnica ctu proceedings 2(1): 178–182, 2015 178 doi: 10.14311/app.2015.02.0178 the hidden population of am cvn binaries in the sloan digital sky survey p. j. carter1, t. r. marsh1, d. steeghs1, e. breedt1, c. m. copperwheat2, b. t. gänsicke1, p. j. groot3, g. nelemans3,4 1department of physics, university of warwick, coventry cv4 7al 2astrophysics research institute, liverpool john moores university, ic2, liverpool science park, 146 brownlow hill, liverpool l3 5rf 3department of astrophysics/imapp, radboud university nijmegen, po box 9010, 6500 gl nijmegen, the netherlands 4institute for astronomy, ku leuven, celestijnenlaan 200d, 3001 leuven, belgium corresponding author: philip.carter@warwick.ac.uk abstract we present results from a spectroscopic survey designed to uncover am canum venaticorum (am cvn) binaries hidden in the photometric database of the sloan digital sky survey (sdss). the discovery of only 7 new am cvns in the observed part of our sample suggests a lower space density than previously predicted. based on the complete g ≤ 19 sample, we calculate an observed space density for am cvns of (5 ± 3) × 10−7 pc−3. we also compare the cataclysmic variables (cvs) discovered via this survey to those found in the sdss spectroscopy, and we discuss sbss 1108+574, an unusually helium-rich cv that has a spectroscopically confirmed orbital period of 55 minutes, well below the cv period minimum (∼80 min). sbss 1108+574 may represent an am cvn forming via the ‘evolved cv’ formation channel. keywords: cataclysmic variables dwarf novae binaries: close stars: individual: sbss 1108+574. 1 introduction the am canum venaticorum (am cvn) binaries are a rare group of hydrogen-deficient, ultra-compact, masstransferring white dwarf binaries. they have orbital periods in the range 5 – 65 minutes, well below the observed period minimum of hydrogen accreting cataclysmic variables (∼80 min, gänsicke et al. 10). three formation channels have been proposed for the am cvn binaries, each characterised by the donor. the donor can be (1) a second, lower mass, white dwarf – these systems have the shortest minimum periods, beginning mass-transfer at periods as short as a few minutes [9, 22]; (2) a semi-degenerate helium star [12, 31]; or (3) an evolved main sequence star [23, 32, 34, 36]. the third channel is known as the ‘evolved cv’ channel, and has generally been considered to be unimportant in comparison to the double white dwarf and helium star channels. the evolution of these systems is thought to be governed by gravitational wave radiation [e.g. 22], causing the mass accretion rate to be a steeply decreasing function of orbital separation. this gives rise to a strong dependence of their observational properties on the orbital period. the shortest period systems are expected to undergo direct impact accretion, and no disc forms [19, 28]. for systems with 10 . porb . 20 min, the accretion disc is in a stable high-state, with spectra showing helium absorption from the optically thick disc [e.g. 21]. am cvns with periods greater than ∼40 min, are thought to be in a stable low-state, the spectra of these systems are characterised by emission lines from the accretion disc [e.g. 30]. in the intermediate period systems (20 . porb . 40 min), an instability occurs in the accretion disc, and their appearance varies between that of the high-state and the low-state systems, similarly to the hydrogen-rich dwarf novae [14, 24]. in the past 10 years, the known population of am cvns has quadrupled, largely as a result of dedicated surveys using the sloan digital sky survey (sdss; york et al. 37), and the palomar transient factory [16, 17]. here we discuss estimates of the am cvn space density, and results from our spectroscopic survey designed to uncover new am cvn binaries amongst colour-selected objects from the sdss photometric database [6, 25, 29]. 2 the serendipitous sdss am cvns the first am cvn binaries were discovered in a variety of different ways, and only 10 were known prior to 178 http://dx.doi.org/10.14311/app.2015.02.0178 the hidden population of am cvn binaries in the sloan digital sky survey the burst of discoveries over the last decade. roelofs et al. [26] and anderson et al. [1, 2] discovered a total of six new am cvn systems in the sdss spectroscopic database via their helium emission dominated spectra. this provided the first sufficiently complete and homogeneous sample of am cvn systems that a study of the population became possible. roelofs et al. [27] used these ‘serendipitous sdss am cvns’ to estimate the observed space density of am cvn binaries – a crucial quantity for calibration of predictions from binary evolution theory. they calculate the completeness of the sdss spectroscopy as a function of colour and galactic latitude, and using an assumed magnitude distribution for am cvns, determine the number and distribution of systems that would be expected from nelemans et al. [20]’s population synthesis. these numbers are then compared to the number of systems found in the sdss, giving a value for the am cvn space density of 1–3 × 10−6 pc−3, an order of magnitude lower than the expected value at the time (2 × 10−5 pc−3, nelemans et al. 20). roelofs et al. [27]’s study also indicated that there should be at least 50 am cvns in total in the sdss photometry, most of them ‘hidden’ due to the absence of spectroscopic data. 3 the search for the hidden population since the completeness of the sdss spectroscopy in the area of colour space occupied by the am cvns is low, and this area is sparsely populated (see fig. 1), roelofs et al. [29] began a dedicated spectroscopic survey of objects in this region, intended to uncover the ∼40 am cvns that were expected [27]. fig. 1 shows the low density of sources in the sdss photometry in the colour region occupied by the am cvn binaries. also shown are the ∼2000 candidates selected from the sdss dr7 database by applying the colour cuts given in roelofs et al. [29]. our targeted region of colour space selects those am cvns with emission line spectra – the longer period systems. as am cvns should spend only a few percent of their lifetime as mass-transferring systems at orbital periods below 30 minutes, our selection should include the vast majority of am cvns in the sdss footprint [6]. this region of colour space is also occupied by db white dwarfs, which are the most significant contaminant in our survey. to date we have taken low-resolution, low signal-tonoise ratio spectra of ∼70% of these candidates, uncovering 30 cvs and 7 new am cvns (see carter et al. 6 for more details). −0.5 −0.4 −0.3 −0.2 −0.1 0.0 0.1 0.2 g − r −0.4 −0.2 0.0 0.2 0.4 u − g figure 1: the greyscale represents the density of sources in the sdss photometric database as a function of colour, to a limiting magnitude g = 20.5 (dereddened). the long period sdss am cvn binaries are indicated by star symbols. the solid line marks the blackbody cooling track, the dotted and dot-dashed lines indicate model cooling sequences for da and db white dwarfs. the dashed lines indicate the colour cuts given by roelofs et al. [29], and the dots indicate the candidates selected. 4 the cv population spectra of the cvs discovered in our survey were presented in carter et al. [6]. we compare this sample to the sdss cv population [33, these proceedings] by plotting the equivalent widths (ews) of their hα and he i 5875 emission lines, see fig. 2. the distributions of the two populations appear similar, with the strongest hα emitters falling outside our survey colour box. there are two obvious outliers visible in fig. 2, css 1122-1110 (sdss j1122-1110; breedt et al. 5) from the sloan sample, and sbss 1108+574 (sdss j1111+5712; carter et al. 7) from our sample, both having much stronger helium emission compared to hydrogen than the majority of cvs. the porb = 59 min cv, v485 cen [3] is also shown in fig. 2 for comparison. followup phase-resolved spectroscopy of sbss 1108+574 reveals an orbital period of 55.3 ± 0.8 minutes [7], consistent with independent photometric and spectroscopically determined periods [13, 18]. this is clear evidence for the evolved nature of the donor, which has been stripped of most of its hydrogen by, or prior to the onset of, mass-transfer. sbss 1108+574 and the similar systems, css 1122179 p. j. carter et al. 1110 [5] and css 1740+4147 (chochol et al., these proceedings), may be evolving along the ‘evolved cv’ evolutionary pathway toward am cvn binaries. 1 :1 2 :1 3 :1 −100 −80 −60 −40 −20 0 ew he i 5875 (å) −350 −300 −250 −200 −150 −100 −50 0 e w h α (å ) css 1122 sdss j0804 sbss 1108 v485 cen figure 2: ew of he i 5875 versus hα for our cv sample (triangles), and the cv population from the sdss (crosses). the star symbols represent the am cvns from our sample, which fall on the ew(hα) = 0 line. also plotted are lines showing hα to he i 5875 ew ratios; cvs with unusually low ew ratios are labelled, v485 cen is shown for comparison. the position of the am cvn sdss j0804+1616 is due to strong he ii 6559 emission. 5 the am cvn space density roelofs et al. [27]’s observed space density, 1–3 × 10−6 pc−3, corresponds to ∼40 am cvns in our survey sample. we have found 7 new am cvns in the ∼70% of the sample that has been observed, suggesting that either there is some bias that we have not considered, or the space density is lower than previously estimated. biases in the survey are discussed in detail by roelofs et al. [29] and carter et al. [6]. they show that due to the distribution in colour space of the spectroscopic observations from the sdss, it is unlikely that a large number of am cvns lie outside the selected colour box. there is a slight bias in our survey observations towards the easier to observe brighter objects, but there is no significant colour bias, and the edges of our colour box are well explored. we therefore consider it likely that a lower space density is required to explain our results. since our survey is essentially complete down to a gband magnitude of 19, we use this smaller sample to recalibrate the population synthesis results. following the method described in roelofs et al. [27] we calculate the expected magnitude distribution of am cvn binaries in the sdss photometric database. the roelofs et al. [27] space density corresponds to 11 am cvns with g ≤ 19. in our essentially complete to g = 19 sample there are 4 known am cvns, scaling the space density to match this number, we obtain a value of (5 ± 3) × 10−7 pc−3 [6]. 6 summary the sdss has significantly increased both the numbers of am cvn stars known, and our understanding of the population. our spectroscopic survey of colourselected objects from the sdss photometric database has uncovered a further seven new am cvns. we have also identified a helium-rich cv, sbss 1108+574, follow-up observations of which reveal an orbital period of 55 minutes, well below the cv period minimum. sbss 1108+574 and similar helium-rich systems may be examples of am cvns forming via the ‘evolved cv’ pathway. discovering only seven new am cvns in the observed sample likely indicates a significantly lower space density than previously predicted. we use the essentially complete, brighter part of our sample to estimate the observed space density to be 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128 [36] tutukov, a. v., et al., 1985, soviet astronomy letters, 11, 52 [37] york, d. g., et al., 2000, aj, 120, 1579 discussion nataly katysheva: do you classify sbss 1108+574 as an am cvn or an su uma star? philip carter: we do not classify sbss 1108+574 as an am cvn because the am cvn classification is classically applied only to objects in which no hydrogen is seen; we decided to refer to it simply as a helium-rich cv. since this system is clearly below the cv period minimum it does not seem entirely appropriate to group it with normal cvs either. whilst it is not certain whether it will become sufficiently hydrogen-depleted to become a ‘true’ am cvn, it is clear that sbss 1108+574 and similar systems lie between the classical definitions of these two classes. koji mukai: is your survey sensitive to am cvn systems with high-state discs, which would have absorption lines, or to direct impact accretors? philip carter: es ceti is a border-line direct impact accretor, showing emission lines [8], and does fall within our colour box. we would have been able to detect systems similar to es cet. those systems showing absorption lines from a high-state disc would appear indistinguishable from db white dwarfs at the low s/n employed in our survey, and so we are not sensitive to them. however, such systems are expected to be a small minority of the total population, and our conclusion about the space density takes this into account. christian knigge: i know the numbers are still small, but have you started to look at the period and magnitude distribution of your systems and checked how they compare to the theoretical predictions? 181 http://dx.doi.org/10.1093/mnras/stt169 http://dx.doi.org/10.1086/427959 http://dx.doi.org/10.1111/j.1365-2966.2009.15126.x http://dx.doi.org/10.1086/509723 http://dx.doi.org/10.1088/0004-637x/739/2/68 http://dx.doi.org/10.1093/mnras/sts672 http://dx.doi.org/10.1046/j.1365-8711.2002.05346.x http://dx.doi.org/10.1093/mnras/271.4.910 http://dx.doi.org/10.1046/j.1365-8711.2003.06380.x http://dx.doi.org/10.1111/j.1365-2966.2011.19924.x http://dx.doi.org/10.1088/0004-637x/708/1/456 http://dx.doi.org/10.1111/j.1365-2966.2005.09186.x http://dx.doi.org/10.1111/j.1365-2966.2007.12451.x http://dx.doi.org/10.1088/2041-8205/711/2/l138 http://dx.doi.org/10.1111/j.1365-2966.2008.14288.x http://dx.doi.org/10.1086/320578 p. j. carter et al. philip carter: we have not yet investigated this in detail, and there are still a few systems for which we do not yet have an orbital period. in a magnitude limited sample, the expected period distribution peaks at ∼50 min, with approximately similar numbers above and below this peak [27]. we have found more systems in the sdss with periods below 50 min, than above, but the numbers are still too small to draw any strong conclusion from this. 182 introduction the serendipitous sdss am cvns the search for the hidden population the cv population the am cvn space density summary 76 acta polytechnica ctu proceedings 2(1): 76–80, 2015 76 doi: 10.14311/app.2015.02.0076 probing the accretion processes in soft x-ray selected polars i. traulsen1, k. reinsch2, a. d. schwope1 1leibniz-institut für astrophysik potsdam (aip), an der sternwarte 16, 14482 potsdam, germany 2institut für astrophysik, georg-august-universität göttingen, friedrich-hund-platz 1, 37077 göttingen, germany corresponding author: itraulsen@aip.de abstract high-energy data of accreting white dwarfs give access to the regime of the primary accretion-induced energy release and the different proposed accretion scenarios. we perform xmm-newton observations of polars selected due to their rosat hardness ratios close to -1.0 and model the emission processes in accretion column and accretion region. our models consider the multi-temperature structure of the emission regions and are mainly determined by mass-flow density, magnetic field strength, and white-dwarf mass. to describe the full spectral energy distribution from infrared to x-rays in a physically consistent way, we include the stellar contributions and establish composite models, which will also be of relevance for future x-ray missions. we confirm the x-ray soft nature of three polars. keywords: cataclysmic variables polars spectroscopy photometry x-rays individual: ai tri, qs tel, rs cae. 1 introduction accretion onto magnetic white dwarfs involves plasma under extreme physical conditions, in particular high temperatures up to millions of kelvin. x-ray observations of the discless am her-type systems (“polars”) provide direct insight into the accretion processes and the opportunity to study related system properties. hard x-ray emission (e > 0.5 kev) arises from the cooling accretion column above the white dwarf and soft (e < 0.5 kev) from the heated accretion region on the white-dwarf surface. model calculations and recent spectral analyses reveal complex structures of the emission regions and a wide range of temperatures and densities. several systems are found at excesses of soft over hard x-ray flux by factors up to 1 000, which can be interpreted as a sign of inhomogeneous accretion. full understanding of the accretion processes and the binary system requires multi-wavelength data, since the different system components dominate the spectral energy distribution (sed) at different wavelengths from infrared up to x-rays. in a campaign of dedicated xmm-newton and optical observations of selected am her-type systems, we spectrophotometrically study the parameters and flux contributions of their components. we concentrate on the conditions in the emission regions in the post-shock accretion column and on the heated white dwarf, flux and luminosity ratios and their strong dependence on the choice of the underlying spectral models. here, we summarize our work on three soft x-ray selected polars and our efforts to establish consistent multi-wavelength models. 2 observed seds starting in 2005, we obtained xmm-newton x-ray and ultraviolet data of ai tri, qs tel, and rs cae (obs. ids 0306840901, 0306841001, 0404710401, 0554740801; traulsen et al., 2010, 2011, 2014), covering one to five orbital cycles per object. on the basis of optical monitoring, the too observations were triggered during high and intermediate high states of accretion. optical photometry and – for qs tel – spectroscopy were performed (quasi)simultaneously. to construct the full long-term seds of the objects from infrared to x-ray wavelengths, we use publicly available archival data, in particular of the wise, 2mass, hst, fuse, and rosat archives. figure 1 shows all data, the main system components being marked according to their approximate flux maxima in the middle panel as follows: a. the secondary star in the ir, b. the cyclotron emission in the ir to optical, c. the accretion stream and d. the white-dwarf primary in the optical to uv, e. the accretion-heated region on the white-dwarf surface in the extreme uv to the soft x-ray regime, and f. the post-shock accretion column in hard x-rays. all panels include both high-state and low-state data, identifiable by their different flux levels. the distinct soft x-ray / euv components at high states are clearly visible, compared for example to the low-state sed of ef eri 76 http://dx.doi.org/10.14311/app.2015.02.0076 probing the accretion processes in soft x-ray selected polars (schwope et al. 2007) or to the serendipitously discovered x-ray hard polar 2xmmp j131223.4+173659 (vogel et al. 2008). at times of high soft x-ray flux, the optical and soft x-ray data show pronounced short-term variability (“flickering”). 10 5 10 4 10 3 10 2 10 1 wavelength [å] 10−6 10−5 10−4 10−3 10−2 10−1 ai tri xmm rosat euve galex fuse iue hst ground 2mass wise 10−6 10−5 10−4 10−3 10−2 10−1 e ne rg y fl ux [ k ev c m − 2 s− 1 ] qs tel a b c d e f 10−6 10−5 10−4 10−3 10−2 10−1 energy [kev] 0.001 0.01 0.1 1.0 10.0 rs cae figure 1: observed seds of three x-ray soft polars from ir to x-rays: our xmm-newton and optical plus archival data at different epochs and accretion states (cf. traulsen et al. 2010, 2011, 2014). the x-ray spectra are unfolded using the best-fit models. 3 multi-wavelength modeling to consistently describe the spectral energy distribution and multi-band light curves, we synthesize spectral models for one uniform set of system parameters and calculate the corresponding light curves. for the different system components, we adopt (referring to the labels used in sect. 2 and fig. 1): a. a phoenix stellar atmosphere model of an m star (hauschildt & baron 1999), b. a cyclotron component of a stratified postshock accretion column (fischer & beuermann 2001), c. a simplified accretion-stream component of a 3d binary model (staude et al. 2001), d. a non-lte whitedwarf atmosphere model (werner & dreizler 1999), e. a singleor multi-temperature black body or hot whitedwarf atmosphere, f. a singleor multi-temperature plasma model of the accretion column. the input parameters of the models are determined from observational data, where possible, and estimated as typical values for primary white dwarf and secondary m star otherwise (e.g. townsley & gaensicke 2009, knigge 2006). in particular, orbital period, inclination, and magnetic field strength are available from optical spectroscopy and polarimetry; parameters of the accretion-induced emission are fitted to the x-ray spectra (cf. sect. 4 and table 1). table 1: parameters of the best fits to the xmmnewton spectra. bolometric rosat flux ratios at fixed temperatures are calculated from (a)schwarz et al. 1998, (b)schwope et al. 1995, (c)burwitz et al. 1996, applying corrections of κbb = 2.5, κbr = 4.8. ai tri qs tel rs cae black body multi-t single-t single-t plasma model multi-t two-t single-t nh,ism [cm −2] ∼ 1020 ∼ 1020 2±1×1019 ktbbody [ev] 44 ± 5 20 ± 4 36 ± 1 ktplasma [kev] 1 .. 20 0.2 .. 4.0 7 ± 3 nh,intr [cm −2] 3±2×1023 1022..1023 ∼ 4×1023 log ṁ [m�/yr] −11.. − 9 ∼ −10 ∼ −10 fbb/fplasma ∼ 3..200 ∼ 10..100 ∼ 11 fbb/fbr, rosat ∼ 170(a) ∼ 80(b) ∼ 90(c) synthetic cyclotron light curves are derived from the phase-dependent model spectra of the accretioncolumn (b) by folding them with the johnson and xmm-newton om filter bandpasses, and white-dwarf and accretion-stream light curves from model components c and d. by comparing them with observational data, we determine their respective phase shifts and intensities. as an example, fig. 2 shows the cyclotron spectra calculated for rs cae, fig. 3 the corresponding synthetic ubv ri light curves. using a consistent set of parameters for all components, we thus establish a physically realistic and consistent multi-wavelength model of the whole binary system, missing only the unknown spectral contribution of the accretion stream. we describe its successful application to the multi-wavelength data of rs cae in traulsen et al. (2014). the most relevant limitation relates to the different observational epochs of the high-state data. while we need low-state spectra in the ir and uv to identify the secondary and (unheated) 77 i. traulsen, k. reinsch, a. d. schwope primary star, all high-state data should be, ideally, observed simultaneously, due to the high variability of the accretion processes. 4 probing the accretion processes in x-rays as described above, x-ray data are not the sole, but the main source of information on the accretion mechanisms in magnetic cvs. they give us access to characteristic determinants of the systems and their evolution like mass accretion rate, component masses, and bolometric fluxes / luminosities, which let us distinguish between accretion scenarios like standing or buried shocks or inhomogeneous accretion. these objectives, however, are limited by the complexity of the puzzle and by the energy resolution and signal-to-noise ratio of currently available x-ray data. several model approaches have been developed, each of them focusing on different aspects, as cropper et al. (1999, mass determination and effective spectral fitting), fischer & beuermann (2001, column structure, sed coverage). mass and flux determination particularly depend on the underlying spectral models (see also cropper et al. 1999). 103 104 105 wavelength [å] 10−21 10−20 10−19 10−18 10−17 10−16 10−15 s im ul at ed f lu x [e rg c m − 2 s− 1 å − 1 ] u b v r i figure 2: spectral models of cyclotron emission for b = 36 mg, ṁ = 0.01, 1.0 g cm−2 s−1, and the estimated accretion geometry of rs cae: orbital mean (red) and phase-resolved from ϕmag = 0.0 (light) to 0.5 (dark). dotted: johnson filter bandpasses, used to derive the light curves in fig. 3. the x-ray soft component is usually modeled by an absorbed black body, which is on the one hand an appropriate approach for ccd data, not resolving the line features. on the other hand, non-lte processes and the metal-richness of the hot photosphere have a non-negligible effect on the spectral continuum in the uv to x-rays (cf. rauch 2003 and references therein). the relevance of consistent non-lte modeling, considering line-blanketing by heavier elements has been demonstrated for non-accreting hot white dwarfs (e.g. traulsen et al. 2005). bolometric fluxes derived from black-body fits to xmm-newton data are lower by factors up to five than from non-lte models, which typically yield lower effective temperatures and hydrogen absorptions. multi-temperature models, designed to reproduce the temperature gradient in the heated accretion region, result in increased bolometric fluxes by at least 50 % with respect to single-temperature models. to illustrate the differences between the models, we simulate pointed observations with the upcoming erosita mission (merloni et al. 2012), which will have a higher effective area at energies between 0.2 and 2 kev than 0.0 0.5 1.0 1.5 magnetic phase 21 20 19 18 17 m ag ni tu de ( v eg a sy st em ) 0.0 0.5 1.0 1.5 x−ray dip phase figure 3: simulated and observed light curves of rs cae. left: ubv ri models (bottom to top), derived from the spectra in fig. 2. right: smarts/b, xmm-newton/u and uvw1. dashed: cyclotron emission, dotted: white dwarf and accretion stream. xmm-newton and rosat. figure 4 shows 50 ks erosita spectra based on fits to the xmm-newton data of ai tri (reduced χ2 between 0.96 and 1.01). for the x-ray hard post-shock spectra, we adopt the radiation-hydrodynamic models by fischer & beuermann (2001). they are valid for shallow accretion columns, both for the shock scenario dominated by bremsstrahlung cooling and the bombardment scenario dominated by cyclotron cooling. to make them available for automated spectral fitting of x-ray ccd and grating data, we incorporate their temperature and density distributions in xspec, using the local mass flow densities ṁ and magnetic field strengths b listed in their paper and white-dwarf masses mwd = 0.6, 0.8, and 1.0 m�. we parameterize the distributions for 30 layers of a stratified column, and add up 30 apec plasma components to the final combined column spectrum. our models include velocity shifts and broadening of the emission lines by stream motion, gravity, orbital motion, and the changing viewing angle. adding a pexmon reflection component (nandra 2007) and multiplying it with the same orbital velocity term, we get a comprehensive description of the phase-dependent emission induced by the accretion column, in particular of the iron lines between 6.4 and 6.9 kev. figure 5 shows composite accretion-column models compared to 78 probing the accretion processes in soft x-ray selected polars a 50 ks synthetic spectrum of an (illustrative) polar with similar fluxes to am her. the models include the same reflection term and different column parameters. the unabsorbed bolometric fluxes of the column models are typically by about 50 % higher than of the corresponding single-temperature fits. the intrinsic absorption and reflection components have a considerable impact on the fluxes (cf. cropper et al. 1999), which increase by factors up to 15 compared to pure plasma models. the soft-to-hard flux ratios, thus, may significantly vary for the same object and observation, depending on the model choice. 0.0 0.5 1.0 1.5 2.0 2.5 n or m al . c ou nt s s− 1 ke v − 1 −5 0 energy [kev] ∆ χ 0.2 0.3 0.4 0.5 + single−t black body + multi−t black body + nlte wd atmosphere figure 4: synthetic spectra of a 50 ks erosita pointed observation, based on different models for the heated white dwarf in the x-ray soft polar ai tri. 3 4 5 6 7 8 9 10 energy [kev] 10 −2 10 −1 1 n or m al iz ed c ou nt s s− 1 ke v − 1 – mwd = 0.6 – 1.0 mo • – b = 100 – 10 mg – m = 0.01 – 100 g cm–2 s–1 ˙ figure 5: simulated 50 ks erosita observation of a polar at ṁ = 1.0 g cm−2 s−1, b = 30 mg, mwd = 0.8 m�, scaled to the xmm-newton spectrum of am her (black: data, yellow: model). each colored line represents one parameter that is varied while the other parameters and the reflection component are fixed (dashed: lowest, solid: highest value). 5 conclusions our xmm-newton observations confirm the soft x-ray excess of the three selected polars and indicate inhomogeneous accretion processes. we develop composite models including the contributions of the stellar atmospheres and the x-ray emitting accretion regions for a physically realistic description of the binary system and the accretion processes. simultaneous multi-λ observations are relevant for a fully consistent sed fitting. the upcoming erosita survey will significantly increase the total number of known systems and of systems for that reliable soft-to-hard ratios can be derived. survey and pointed observations will enable us to better distinguish between different models of the accretion processes, refine them, and push our knowledge about the physical properties of the x-ray emission regions of polars. acknowledgement our research was supported by dlr under grant numbers 50 or 0501, 50 or 0807, and 50 or 1011. references [1] burwitz, v. et al.: 1996, a&a 305, 507 [2] cropper, m. et al.: 1999, mnras 306, 684 doi:10.1046/j.1365-8711.1999.02570.x [3] fischer, a., beuermann, k.: 2001, a&a 373, 211 [4] hauschildt, p. h., baron, e.: 1999, j. comp. appl. math. 109, 41 doi:10.1016/s0377-0427(99)00153-3 [5] knigge, c.: 2006, mnras 373, 484 doi:10.1111/j.1365-2966.2006.11096.x [6] merloni, a. et al.: 2012, arxiv:1209.3114 [7] nandra, k. et al.: 2007, mnras 382, 194 doi:10.1111/j.1365-2966.2007.12331.x [8] rauch, t.: 2003, a&a 403, 709 [9] schwarz, r. et al.: 1998, a&a 338, 465 [10] schwope, a. d. et al.: 1995, a&a 293, 764 [11] schwope, a. d. et al.: 2007, a&a 469, 1027 [12] staude, a., schwope, a. d., schwarz, r.: 2001, a&a 374, 588 [13] townsley, d. m., gänsicke, b. t.: 2009, a&a 693, 1007 [14] traulsen, i. et al.: 2005 in 14th european workshop on white dwarfs, d. koester & s. moehler (eds.), asp conf. ser. 334, 325 [15] traulsen, i. et al.: 2010, a&a 516, a76 [16] traulsen, i. et al.: 2011, a&a 529, a116 79 http://dx.doi.org/10.1046/j.1365-8711.1999.02570.x http://dx.doi.org/10.1016/s0377-0427(99)00153-3 http://dx.doi.org/10.1111/j.1365-2966.2006.11096.x http://dx.doi.org/10.1111/j.1365-2966.2007.12331.x i. traulsen, k. reinsch, a. d. schwope [17] traulsen, i. et al.: 2014, a&a 562, a42 [18] vogel, j. et al.: 2008, a&a 485, 787 [19] werner, k., dreizler, s.: 1999, j. comp. appl. math. 109, 65 discussion christian knigge: your data on ai tri seems to show that the hard x-ray flux actually decreases as the accretion rate (and soft x-rays) increase. what is the interpretation of this? klaus reinsch: it is related to the response of the accretion shock height to the changing specific accretion rate. in addition, a higher fraction of inhomogeneous accretion events can suppress x-ray hard emission. the first observation of ai tri during an extremely soft state does not cover a full binary orbit. 80 introduction observed seds multi-wavelength modeling probing the accretion processes in x-rays conclusions 257 acta polytechnica ctu proceedings 2(1): 257–260, 2015 257 doi: 10.14311/app.2015.02.0257 optical low resolution spectroscopic observations of t pyx during the early phase of 2011 outburst a. arai1,2, m. isogai2,3, m. yamanaka4, h. akitaya5, m. uemura5 1center for astronomy, university of hyogo 2koyama astronomical observatory, kyoto sangyo university 3national astronomical observatory of japan 4kwasan observatory, kyoto university 5higashi-hiroshima astronomical observatory, hiroshima university corresponding author: arai@nhao.jp abstract we report on the results of our low resolution spectroscopic observations during the 2011 outburst of the recurrent nova t pyx. our observations were performed from 0.19 days to 34 days after the eruption discovered by m. linnolt. we found wolf-rayet like features in our spectrum during the initial rising phase on t = 0.19 d. following spectral developments are consistent with previous works. we discuss that the early phase of t pyx is divided into three stages, a short lived wr-like stage, he/n stage and fe ii stage. keywords: cataclysmic variables optical spectroscopy individual: t pyx. 1 introduction the initial rising phase of novae is a very intriguing phase in the point of view of thermonuclear runaway and development of spectral features. generally, observations are very poor in the initial rising phase, because it is very rare that novae are discovered during their initial rising phase. especially, spectroscopic observations lack at these very early epochs. t pyx is one of the recurrent novae. its previous outbursts were observed in 1890, 1902, 1920, 1944, and 1967 (schaefer et al. 2010). low resolution spectroscopic observations in past outbursts were performed around their visual maximum (e.g., adams and joy 1920, joy 1945, and catchpole et al. 1969). recurrent novae (rne) are semi-detached binary systems consisting of a massive white dwarf with accretion material from a companion star (at typical mass transfer rate ' 10−7 m� yr−1). they are the prime candidates for being type ia supernovae (sne) progenitor. recurrent novae are classified into three types, the u sco-, the rs ophand the t pyx-types, based on the light curves and spectroscopic features (anupama et al. 2008). but, the light curve of t pyx in the early phase is more similar to slow rising classical novae, than the u scoor the rs oph-type rne. in past outbursts, t pyx showed a rapid rising to v ∼ 8, followed by a pre-maximum halt and a further slow rising to the maximum light (v ∼ 6.5) for ∼ 20 days with variations of ∆v ∼ 0.2. the sixth outburst of t pyx was discovered at visual magnitude 13.0 on ut 2011 april 14.2931 (= jd 2455665.7931, t = 0 d) by m. linnolt (waagan et al. 2011). after the discovery, many observers carried out follow up observations. the study in the early spectroscopic observations and photometry revealed that t pyx showed rare hybrid spectral classification types (williams 1992), he/n fe ii he/n transition (e.g. izzo et al. 2012, imamura and tanabe 2012, surina et al. 2014). following prompt target of opportunity observations were also performed in x-ray and radio just after the discovery of the outburst (e.g., kuulkers et al. 2011, nordsieck and shara 2011, chomiuk et al. 2011). here, we report our spectroscopic results of t pyx from the very early phase. in this paper, we focused on the wolf-rayet feature in the initial rising phase on t = 0.19 d (2011 april 14.48) and following developments of spectral features from t = 2 d to t = 34 d. 2 observations our low dispersion optical spectroscopic observations of t pyx were performed during 13 nights from 2011 april 14 to may 18 at two sites. one site is koyama astronomical observatory (kyoto sangyo university). we used the araki telescope with losa/f2 spectrograph (shinnaka et al. 2013). the wavelength coverage 257 http://dx.doi.org/10.14311/app.2015.02.0257 a. arai et al. is 4000–8000 å and the spectral resolution is r ∼ 580 at 6563 å. another one is higashi-hiroshima astronomical observatory (hiroshima university). we use the kanata telescope with howpol (kawabata et al. 2008). the wavelength coverage is 4200–9000 å and the spectral resolution of r ∼ 400 at 6563 å. the v-band light curve and epochs of our spectroscopic observations are shown in the panel (a) of figure 1. -800 -700 -600 -500 -400 -300 -200 -100 0 0 5 10 15 20 25 30 35 40 e .w . [a n g st ro m s] day from the discovery (c) 400 600 800 1000 1200 1400 1600 1800 f w h m [ km s -1 ] (b) hα hβ oi 7774 feii 5169 hei 5876 5 6 7 8 9 10 11 12 v -b a n d m a g n itu d e s (a) figure 1: (a) the v-band light curve of the early phase of t pyx 2011 quoted from aavso database. short tics denote epochs of our observations. (b) the fwhms of prominent emission lines. (c) the equivalent widths for the same lines in the panel (b). 3 results and discussion 3.1 wolf-rayet like features in the initial rising phase on t = 0.19 d figure 2 shows our first spectrum on ut 2011 april 14.48 (t = 0.19 d), just 4.4 hours after the discovery by m. linnolt. the spectrum exhibits many highly excited emission lines, he ii, c iv, n iii, n iv, on this spectrum. the spectral features are reminiscent of wn type wolf-rayet (wr) stars (crowther 2007 and references therein). we identified these emission lines with reference to spectral atlases of wolf-rayet stars (hamann 1995, conti 1990). the strong emission at c vi (λ5802) is, however, observed in wc type rather than wn type. in wolf-rayet stars, emission lines of highly excitation levels for c, n or o are often observed as strong broaded emission lines (fwhm = 800 2000 km s−1). in generally, the stellar temperature for wr stars is around 5 × 104 – 105 k, and wr winds are accelerated by line driven pressure (crowther 2007). in the ordinary picture of novae, during the initial rising phase, shells / a photosphere is expanding. absorption components are expected to be detected for the balmer and fe ii, o i, or he i transition. t pyx showed, however, no significant p-cygni profiles at the expanding velocities of highly excited emission lines (fwhm ∼ 1100 km s −1) on t = 0.19 d. this implies that the envelope of wd would be optically thin on the moment, that is, the optically thick wind (kato 1994) would be still weak on that time. the t pyx binary system has an orbital period of 1.83 h and a mass ratio of 0.2±0.03 (uthas et al. 2010). such compact binary system has a dwarf companion, suggesting that the amount of surrounding gas supplied by the companion star of t pyx is much less than the rs oph-type rne, which show highly excited emission by shock (e.g., [fe vii]). hence, the wolf-rayet like spectral features of t pyx would be originated in the nova envelope, not by shocks with the circumbinary gas. at this time, wolf-rayet like spectral features are reported for a few novae. in the pu vul (1979), wn type features are discovered in optical and uv during its decline phase (tomov et al. 1991, sion et al. 1993). in the spectra of ag peg, rr tel and rt ser, wr-like emission lines appeared after the visual maximum (kenyon 1986, tomov et al. 1991 and references therein). our result is the first case of wr-like spectral features in the rising phase of novae. 3.2 developments after t = 2 d such wr-like spectral features disappeared in our spectra from t = 2 d as shown in figure 3. no wr-like lines were reported after t = 0.8 – 1 d (shore et al. 2011, izzo et al., 2012, surina et al. 2014). the brightness of t pyx on t = 2 d was v = 8, in the pre-maximum halt as shown in the panel (a) of figure 1. on t = 2 d, hα, hβ, o i, he i and n ii have been strengthen, while emission lines of he ii, c iv and n iv disappeared. this result indicates that the spectral type of t pyx had changed from wr-like type to he/n type. the expanding velocity estimated from blue-shifted absorption minimum of p-cygni profiles of hβ is about 1400 km s−1. 258 optical low resolution spectroscopic observations of t pyx during the early phase of 2011 outburst 1.0 2.0 3.0 4.0 5.0 6.0 4000 4500 5000 5500 6000 6500 7000 7500 n o rm a liz e d f lu x wavelength (å) h δ/ h e i i + n i v h e i i 4 1 9 9 h γ n i ii 4 5 1 5 h e i i 4 5 4 1 + c i v n v 4 6 0 3 + 4 6 1 8 n i ii 4 6 5 0 h e i i 4 6 8 6 h β o v 4 9 4 0 o r h e i 4 9 2 2 , f e i i 4 9 2 4 , f e v ii 4 9 4 0 f e i i 5 5 8 1 + 5 5 8 2 + 5 5 9 0 c i v 5 8 0 2 h α h e i i 6 6 8 3 n i v 7 1 0 9 − 7 1 2 3 h e i i 7 1 7 7 ⊕ ⊕ figure 2: first spectrum of t pyx on t = 0.19 d. the highly excited emissions were detected likely to wolf-rayet stars. cross circles denote positions of telluric absorption bands. 0 5 10 15 20 25 4000 4500 5000 5500 6000 6500 7000 7500 8000 n o rm a liz e d f lu x wavelength (å) h α h β h γ h δ f ei i h ei h ei /n ii h ei h ei h ei o i t=0.19 kao t=2 hhao t=3 kao, hhao t=5 kao t=6 kao t=7 kao t=9 hhao t=11 kao t=12 hhao t=18 kao t=29 kao, hhao t=33 kao t=34 kao figure 3: low resolution spectra of t pyx from t = 0.19 – 34 d. observations were carried out at koyama astronomical observatory (kao) and higashi-hiroshima astronomical observatory (hhao). we identified the balmer, he i, n ii, fe ii multiplets, o i, na i lines. the developments of fwhms and equivalent widths of emission components of these lines are shown in the panel (b) and (c) of figure 1, respectively. the fwhm of hα (and also hβ) measured in our spectra (the panel (b) of figure 1) present a rapid increase from 1100 km s−1 to 1600 km s−1 during t = 0.19 – 2 d. surina et al. (2014) reported faster velocity (vej = 4000 km s −1) on t = 0.8 d. after this event, the fwhm of hα decreased gradually to ∼ 1000 km s−1 until t = 12 d in parallel with the increasing of their equivalent widths from -70 å to -320 å, as shown in the panel(c) of figure 1. fwhms of other emission lines also showed similar trend, and their equivalent widths decreased to the minimum around t = 20 d (e.w. of hα = -200 å). from t = 12 d to t = 34 d, fwhms of all lines increased again, from ∼ 1000 km s−1 to ∼ 1300 km s−1 for hα, hβ and o i (λ7774). hβ, n ii (λ5679), o i (λ7774) are accompanied with p-cygni profiles clearly. t pyx showed a he/n-type spectra from t = 2 d to t = 6 d, during pre-maximum halt. decrease rates of equivalent widths of he i and n ii are similar to the one observed in the balmer lines at the same epoch. on the other hand, fe ii emission lines emerged aound t = 7 d as seen in the panel (c) in figure 1. the equivalent widths of fe ii multiplets(42, 49) were growing gradually from t = 7 d to t = 34 d. in this epoch, increasing trend of fwhm in fe ii and o i were also observed as seen in hα and hβ. on t = 34 d, the equivalent width of fe ii and o i were slightly decreased due to the growth of continuum light in this epoch. these results indicate that the strong optically thick wind were gradually developing and fe ii type features replace a he/n features around t = 7 d, consisting with surina et al. (2014). after the entering into the fe ii type, hβ, n ii (λ5679), o 259 a. arai et al. i (λ7774) are accompanied with clear p-cygni profiles. these spectral developments are also reported and discussed in detail in shore et al. (2011), izzo et al. (2012), imamura and tanabe (2012), and surina et al. (2014). as pointed out by imamura and tanabe (2012), a similar hybrid behavior during the pre-maximum halt phase was also observed in only a slow developing nova, v5558 sgr (tanaka et al. 2011). t pyx reveals intriguing spectral transition of spectral type of nova at more earlier stage than that observed in v5558 sgr. the hybrid (he/n fe ii) transition would be a common behavior in slow rising novae. very likely, the early observed wr-like features may be also a common evolution just after the explosion for slow rising novae. based on our results and combination with previous studies, we summarize that the development during the very early phase of the 2011 outburst of t pyx is composed of three stages: (1) the wr-like stage, which shows short lived (duration < 1 d) wr-like spectral line features, (2) the he/n stage (several days), and (3) the fe ii-type stage. acknowledgement we thank taichi kato (kyoto university) for valuable comments on our observational data. this research was supported(, in part,) by a grant from the hayakawa satio fund awarded by the astronomical society of japan, and optical & near-infrared astronomy inter-university cooperation program, supported by the mext of japan. references [1] adams, w. s. and joy, a. h. 1922, publications of the american astronomical society, 4, 139 [2] anupama, g. c., evans, a., bode, m. f., o’brien, t. j. and darnley, m. j. 2008, astronomical society of the pacific conference series, 40, 31 [3] catchpole, r. m. 1969, mnras, 142, 119 doi:10.1093/mnras/142.1.119 [4] chomiuk, l., et al. 2011, the astronomer’s telegram, 3318 [5] conti, p. s. and massey, p. and vreux, j.-m. 1990, apj, 354, 359 doi:10.1086/168694 [6] crowther, p. a. 2007, ara&a, 45, 177 doi:10.1146/annurev.astro.45.051806.110615 [7] hamann, w.-r. and koesterke, l. and wessolowski, u. 1995, a&as, 113, 459 [8] imamura, k. and tanabe, k. 2012, pasj, 64, l9 [9] izzo, l., et al. 2012, mmsai, 83, 830 [10] joy, a. h. 1945, pasp, 57, 171 doi:10.1086/125711 [11] kato, m. and hachisu, i. 1994, apj, 437, 802 doi:10.1086/175041 [12] kawabata, k. s., et al. 2008, spie, 7014, 4 [13] nordsieck, k. h. and shara, m. 2011, the astronomer’s telegram, 3289, 1 [14] kenyon, s. j. 1986, aj, 91, 563 [15] kuulkers, e, et al. 2011, the astronomer’s telegram, 3285, 1 [16] schaefer, b. e., pagnotta, a. and shara, m. m. 2010, apj, 708, 381 doi:10.1088/0004-637x/708/1/381 [17] shinnaka, y., et al. 2013, icarus, 222, 734 doi:10.1016/j.icarus.2012.08.001 [18] shore, s. n., augusteijn, t., ederoclite, a. and uthas, h. 2011, a&a, 533, 8 [19] sion, m. edward et al. 1993, aj, 106, 2118 [20] surina, f., et al. 2014, aj, 147, 107 [21] tanaka, j., et al. 2011, pasj, 63, 911 [22] tomov, t., et al. 1991, mnras, 252, 31 doi:10.1093/mnras/252.1.31p [23] uthas, h. and knigge, c. and steeghs, d. 2010, mnras, 409, 237 doi:10.1111/j.1365-2966.2010.17046.x [24] waagan, e., et al. 2011, cbet, 2700, 1 [25] williams, r. e. 1992, aj, 104, 725 260 http://dx.doi.org/10.1093/mnras/142.1.119 http://dx.doi.org/10.1086/168694 http://dx.doi.org/10.1146/annurev.astro.45.051806.110615 http://dx.doi.org/10.1086/125711 http://dx.doi.org/10.1086/175041 http://dx.doi.org/10.1088/0004-637x/708/1/381 http://dx.doi.org/10.1016/j.icarus.2012.08.001 http://dx.doi.org/10.1093/mnras/252.1.31p http://dx.doi.org/10.1111/j.1365-2966.2010.17046.x introduction observations results and discussion wolf-rayet like features in the initial rising phase on t = 0.19d developments after t = 2d 79 acta polytechnica ctu proceedings 1(1): 79–83, 2014 79 doi: 10.14311/app.2014.01.0079 forty years of x-raying narrow-line seyfert 1 galaxies thomas boller1 1max-planck-institut für extraterrestrische physik, garching, germany corresponding author: bol@mpe.mpg.de abstract forty years after the discovery of narrow-line seyfert 1 galaxies (nls1s), and 20 years since the discovery of the remarkable ultrasoft soft x-ray emissions of nls1s, strategic publications improved the understanding of the seyfert phenomenon more generally. new theoretical models emerged from the observations and stimulated the discussions on the innermost regions of agn. nls1s are an amazing class of agn for x-ray, optical and multiwavelength science. keywords: x-rays: narrow-line seyfert 1 galaxies. 1 introduction type 1 agn are divided into broad line seyfert 1 galaxies (bls1s) and narrow-line seyfert 1 galaxies (nls1s). bls1s are associated with the name of carl seyfert (∗11.2.1911, cleveland, ohio, usa+13.6.1960, nesville, tenneesee, usa). seyfert (1943) discovered that the fwhm of the optical emission lines reaches values up to about 10000 km s−1. such high values were thought to originate from the broad-line region in the gravitational potential of supermassive black hole. the fwhm values of bls1s are much higher than the velocity dispersions found from the rotation curves of normal galaxies, which reach only a few 100 km s−1. observations of such high velocities in bls1s suggested in addition that the observer has a direct view to the central regions of active galaxies. in 1970, fritz zwicky (∗14.2.1898, bulgaria,+ 8.2.1974, pasadena, califormia, usa) made the fundamental discovery that type 1 agn exhibit fwhm of the optical permitted lines much less than that what was observed in bls1s, with fwhm values down to 500 km s−1 (zwicky 1970). these objects were sometimes misclassified with normal galaxies, however the presence of very strong fe ii multiplet emission definitely revealed the type 1 agn nature of these sources, as fe ii multiplet emission is only occurring in high density regions with densities larger than 109 cm−3. such densities are only present in the blr clouds or the accretions disc in agns. 2 historical review 1970, zwicky the first note on a new class of type 1 agn with unusually narrow permitted lines was made by zwicky (1970). 1978, davidson and kinman davidson and kinman (1978) investigated the optical spectrum of mrk 359. they pointed out that the object has unusually narrow permitted optical lines and that ’this object merits further observations‘. figure 1: fwhm value of the hβ line versus the rosat photon index for broad line seyfert 1 galaxies and narrow-line seyfert 1 galaxies. below 2000 km s−1 the photon index reach values up to about 5. 79 http://dx.doi.org/10.14311/app.2014.01.0079 thomas boller 1985, osterbrock and pogge osterbrock and pogge (1985) came up with the first definition for nls1s. they defined nls1s as follows: (i) the permitted optical lines are only slightly broader than the forbidden lines; (ii) strong fe ii multiplet emission, centered between 4400 and 4500 å; (iii) a ratio of the o iii line to hβ line with less than 3; (iv) having fwhm of the hβ line between 500 and 2000 km s−1 (added by goodrich 1989). 1992, puchnarewicz based on einstein ipc data (puchnarewicz 1992) reported for the first time on remarkable ultra-soft x-ray spectra of nls1s. the sample size of nls1 with steep soft x-ray spectra reached a number of 53. 1996, boller, brandt, fink in 1996, we (boller, brandt, fink, 1996) published a relation between the fwhm value of the hβ line and the rosat 0.1-2.4 kev photon index (c.f. fig. 1). while bls1 have their photon indices confined to a fairly narrow range around about 2.3, the rosat photon index rises steeply below fwhm values of 2000 km s−1. this plot demonstrates that the emission within a few rg around the black hole determines the velocity distribution at much larger scales in the blr. the main reasons for the steep x-ray spectra of nls1s are the low masses of the black hole and the very high accretion rates (boller et al. 1996). there appears also a zone of avoidance (fwhm values larger than 2000 km s−1 and photon indices larger than 3), where so seyfert galaxies have been observed so far. 1997, boller, brandt, fabian, fink figure 2: 30 day rosat monitoring of the nls1 galaxy iras 13224-3809 taken in 1997. persistent, giant and rapid amplitude variations have been detected with the maximum amplitude variation with a factor of about 57. in 1997 we have performed a 30 day rosat monitoring observation of the nls1 galaxy iras 13224-3809 (boller et al. 1997). giant, persistent and rapid amplitude variations have been discovered with a maximum amplitude variation with a factor of 57 within about 1200 seconds. the peak in luminosity corresponds to about 1044 erg s−1 within that time scale. the most plausible explanation for these extreme x-ray variability is flux boosting of hot spots orbiting the central black hole. as the flux boosting factor scales with the doppler-factor to the power of 3+γ, and as the photon index of the object is very steep with 4.4, such extreme variability is a natural consequence of flux boosting in steep spectrum nls1s (fig. 2). figure 3: sharp spectral cut-off detected in 1h0707495. a power-law plus absorption edge model, a partial covering model and an ionized reflection disk model are shown with their corresponding residua. 2002, boller, fabian, sunyaev, trümper in 2002 we (boller et al., 2002) have discovered the first sharp spectral drop at around 7 kev without any noticeable fe k emission. the spectral drop is of the order of about 2. two explanations have been discussed in this paper, (i) a partial covering scenario, and (ii) a reflection dominated spectrum. while the partial covering model appears to be ruled out in subsequent analysis, the reflection model of ross and fabian (2005) might explain this type of new spectral cut-offs, now also discovered in other nls1s. 2003, pounds high velocity outflows with substantial fractions of the speed of light have been detected by (pounds et al. 2003). in the nls1 galaxy pg 1211+143 outflow velocities up to 0.2 c have been seen. this discovery fits into the general scheme of understanding nls1 as super-eddington accreeting sources, which naturally leads to substantial mass outflows. 80 forty years of x-raying narrow-line seyfert 1 galaxies 2004, miniutti, fabian triggered by the detection of sharp spectral drops in the high energy spectra of nls1s, a new spectral model has been developed, named as the ’light bending model’ by miniutti and fabian (2004). in this model a hot and compact region, emitting a power law spectrum, is located very close to the central black hole and changes its scale height above the black hole. the compact source illuminates the accretion disc and gives rise to reflection spectrum. when the scale height of the compact object is small (2 rg) then most of the power law photons from the compact object are bent towards the black hole and the power law contribution for the external observer is small. the spectrum is in this case reflection dominated and especially the strongest x-ray emission lines, the fe k and fe l lines become visible (fabian et al. 2009). when the source height is increasing, a smaller amount of photons is lost into the black hole and the powerlaw component increases in strength for the external observer. while the power law component varies by a factor of about 70, the reflection component is less variable with a factor of about 4. this model can explain the missing relation between the illumination strength of the disc and the strength of the fe k line as well as most of the spectral and timing properties we have observed from the innermost regions of agns. 2007, komossa in 2006, komossa (2006) carried out a systematic search for radio-loud nls1s. the analysis confirmed that most nls1s are radio-quiet objects and that only 2 per cent can be considered as radio-loud with an r value greater than 100. the physical reasons for the radio-loudness of some nls1s are still under discussions. 2008, foschini; malizia foschini (2008) and malizia (2008) first reported on the detection of nls1s at hard (2-20 kev) x-rays. this was something unexpected given the steep x-ray spectra of these objects. 2009, fabian a very important discovery, strongly supporting the light bending model from miniutti and fabian (2004), was the detection of a 30 second reverberation lag between the direct hard power-law emission and the soft fe l line reflection component, with the power-law emission leading the soft x-ray emission (fabian et al. 2009). the light bending model predicted that the external observer first sees the direct power-law emission from a compact and hot region sitting very close to the central black hole and after that the reflection component, which has a longer light travel time. in addition, both the fe k and the fe l line were seen for the first time in a agn. the lines are strongly relativistic and the intensity of the lines are in the ratio of 20 to 1, in excellent agreement with atomic physics. 2012, foschini the first detection of gamma-ray emission from a nls1 galaxy was reported by foschini et al. (2012, and references therein). a more than three years monitoring programme ranging from the radio to gamma-rays on the nls1s pmn j0948+0022 revealed the detection of emission above 100 mev. the observations revealed the presence of a powerful relativistic jet with isotropic emission of about 1048erg s−1. a complete understanding of the radio and gamma-ray emission has still not emerged from the few available data. 3 upcoming nls1 erosita science 3.1 survey science erosita will carry out a 4 years all-sky survey in the 0.3 to 10 kev band. the expected number of new nls1s which will be detected is 100000, exceeding the presently known number of nls1s by orders of magnitude. these objects can easily disentangled from other objects, via their distribution in the fwhm photon index plane (c.f. fig. 1). as most of the source will have only moderate count rate statistics during the survey observations, the survey science will be restricted to statistical analysis on different science cases discussed below. • eigenvector 1 analysis brandt and boller (1999) have carried out an eigenvector 1 analysis of broad and narrow-line seyfert 1 galaxies. we found that nls1s lie towards one extreme of the primary eigenvector, and that steep x-ray spectra, narrow optical permitted lines, extreme x-ray variability, week [o iii] emission, strong fe ii multiplet emission are correlated. these type of analysis will be applied to the erosita nls1s with much higher data statistics, probing the physical mechanism lying behind these correlations with much better statistics. • super-eddington accretion and comptonization most nls1 galaxies are accreeting above the eddington limit. the relation between supereddington accretion, outflows, steepness of the x-ray spectra and comptonization has been summarized by boller (2011). the basic idea is that super-eddington accretion results into a strong uv luminosity and a stronger and to higher energies shifted planck-emission from the accretion 81 thomas boller disc. the stronger the uv disc luminosity is and the larger the solid angle subtended by the reflector is, the greater is the cooling by the seed photons incident on the plasma, the lower is the plasma electron temperature, and the steeper are the x-ray photon indices, consistent with the observations. again, with the much larger number of new nls1s detected in the erosita survey observations, a better physical understanding of the relation between all these properties will emerge. other science cases which will be addressed during the erosita survey observations will be, (i) search for extreme x-ray variability and nonlinear x-ray variability, (ii) αox science and the global x-ray baldwin effect, and (iii) multiwavelength source population properties. 3.2 erosita science with pointed observations after the 4-years period of survey observations, pointed observations will be carried out with erosita, allowing for much higher count rate statistics. some of the basic science drivers for the nls1 research will be: • relativistic iron k and l lines and reverberation lag physics erosita will offer a unique possibility to further study the behavior of matter under strong gravity on a much larger sample as presently known. • erosita black hole growth studies nls1 galaxies with their proven high accretion rates are the objects with the highest black hole growth rates and are therefore ideally suited for such studies. • accretion disc physics and nls1 science erosita will open a new window on the study of relativistic disc reflection. strong x-ray reflection is a natural consequence for the existence of a compact corona close to the central black hole system. erosita will offer much more nls1 and agn science. among them are: (i) high velocity outflows and feedback processes, (ii) strong gravity in the highcurvature regime and gr tests (boller and müller, 2013), (iii) intensive starbursts, super-solar metallicities, and relation to agn accretion, (iv) extending the luminosity function of nls1s to higher redshifts, (v) models for the black hole regions of nls1s, (vi) the origin of the extreme soft x-ray excesses, (vii) thin and thick accretion discs in nls1s, (viii) the accretion disc coronae of nls1s, (ix) the origin of the extreme x-ray variability in nls1, (x) interpretation of line and continuum correlations, and (xi) models for the broad line region and the x-ray and uv absorption. acknowledgement tb is grateful to luigi foschini for critical reading of the manusscript and franco giovannelli for organizing and leading the workshop. references [1] seyfert, c: 1943, apj 97, 28. doi:10.1086/144488 [2] zwicky, f.: 1970, pasp 82, 93. doi:10.1086/128889 [3] davidson, k., kinman, t.d.: 1989, apj 225, 776. [4] osterbrock, d.e., pogge, r.w.: 1985, apj 297, 166. [5] puchnarewicz, l.: 1992, a&a 256, 589. [6] boller, th., brandt, w.n., fink, h.: 1996, a&a 305, 53. [7] boller, th., brandt, w.n., fabian, a.c. et al.: 1997, mnras 289, 393. [8] boller, th., fabian, a.c., sunyaev, r., trümper et al.: 2002, mnras 329, 1 doi:10.1046/j.1365-8711.2002.05040.x [9] goodrich, r. w.: 1989, apj 342, 224. [10] ross, r.r., fabian, a.c.: 2005, mnras 358, 211. doi:10.1111/j.1365-2966.2005.08797.x [11] pounds, k.: 2003, mnras, 346, 1025. doi:10.1111/j.1365-2966.2003.07164.x [12] miniutti, g., fabian, a.c.: 2004, mnras 349, 1435. doi:10.1111/j.1365-2966.2004.07611.x [13] foschini, l., maraschi, l., ghisellini, g. et al: 37th cospar scientific assembly. held 13-20 july 2008, in montral, canada., p.917 [14] malizia et al.: 2008, mnras, 389. [15] fabian, a.c., et al.: 2009, nature, 459, 540. doi:10.1038/nature08007 [16] komossa, s. et al.: 2006, aj, 132, 531. [17] foschini, l. et al.,: 2012, a&a 548, 106. [18] brandt, w.n., boller, th.,: 1999, aspc 175, 265. [19] boller, th.: 2011, the x-ray universe 2011, presentations of the conference held in berlin, germany, 27-30 june 2011 82 http://dx.doi.org/10.1086/144488 http://dx.doi.org/10.1086/128889 http://dx.doi.org/10.1046/j.1365-8711.2002.05040.x http://dx.doi.org/10.1111/j.1365-2966.2005.08797.x http://dx.doi.org/10.1111/j.1365-2966.2003.07164.x http://dx.doi.org/10.1111/j.1365-2966.2004.07611.x http://dx.doi.org/10.1038/nature08007 forty years of x-raying narrow-line seyfert 1 galaxies [20] boller, th., müller, a., fias interdisciplinary science series, isbn 978-3-319-00046-6, springer international publishing switzerland 2013, p. 293. discussion matteo guainazzi’s comment: do you believe that measurements of black hole masses in nls1s are nowadays robust? thomas boller: after the discovery of the extreme x-ray properties of nls1s all reverberation measurements have shown that nls1s lie at the lower mass end of the mass distribution in agn. especially the papers by b. peterson are extremely helpful in that matter. 83 introduction historical review upcoming nls1 erosita science survey science erosita science with pointed observations acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0027 acta polytechnica ctu proceedings 4:27–32, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app self-absorption corrections based on monte carlo simulations kamila johnová faculty of nuclear science and physical engineering, czech technical university in prague, břehová 7, 115 19 prague, czech republic correspondence: kamila.johnova@fjfi.cvut.cz abstract. the main aim of this article is to demonstrate how monte carlo simulations are implemented in our gamma spectrometry laboratory at the department of dosimetry and application of ionizing radiation in order to calculate the self-absorption within the samples. a model of real hpge detector created for mcnp simulations is presented in this paper. all of the possible parameters, which may influence the self-absorption, are at first discussed theoretically and lately described using the calculated results. keywords: mcnp, self-absorption, gamma spectrometry. 1. introduction gamma spectrometry is powerful non-destructive method useful for samples containing gamma ray emitting radionuclides (spontaneous or induced gamma ray activity). due to the principles of gamma ray emission, energy of photons emitted during radioactive decay is specific for each radionuclide. this means that with a proper detection technique it is possible to identify what radionuclide is present in the sample. in addition, with a suitable efficiency calibration the amount of radionuclide in the sample can be estimated. gamma spectrometry can be applied in many scientific fields, like contamination studies, natural radioactivity, activation analysis, astroand high energy physics or geology studies. different applications require a different detection techniques and data analyses. this paper discuss a laboratory gamma spectrometry for natural samples and samples of natural origin. 1.1. interactions of gamma radiation with matter for the issue discussed in this paper three type of photon interactions are the most important: photoelectric absorption, compton scattering and pair production. the probability that a certain photon of a given energy will interact in the matter is expressed by its cross section. the cross section can sometimes be expressed in special unit barn = 10−28 m2. the sum of cross sections for all possible interaction is denoted as total cross section σt . the cross section is closely connected to the attenuation of photons in the matter. mono-energetic photon beam is along a path of length d attenuated exponentially, which is expressed by equation i = i0e−µd, where i0 is a number of photons in non-attenuated beam, i is a number of photons at the end of path d and µ is a linear attenuation coefficient. µ can be expressed in the terms of cross sections for individual interactions (τ-photoelectric absorption, σ-compton scattering, κ-pair production, σrs-elastic scattering) or in the term of total cross section σt µt ρ = na a (τ + σ + κ + σrs) = na a σt , where µt means total attenuation coefficient (over all interactions), ρ is the density of material, na the avogadro constant and a is an average atomic mass. the expression µt ρ is called total mass attenuation coefficient. 1.2. self-absorption coefficient the previous text suggested that there has to be a certain level of self-absorption within the sample itself. based of the previous informations one can correctly assume that this effect depends on composition of the sample, measurement geometry, density of the material and energy of photons. in general, there are 3 ways how to estimate the self-absorption correction. estimation based on mass attenuation coefficients. several spectrum analysis programs give the user this option, where the library of mass attenuation coefficients is provided. this method is useful over the energy range, where the compton scattering is the dominant interaction. according to the fact, that mass attenuation due to compton scattering is almost independent of an atomic number, reasonable self-absorption correction can be made even for the sample of barely known composition. empirical estimation. this procedure can sometimes be useful for laboratories providing routine measurements. it is specially useful, when the sample of unknown composition and containing lowenergy emitting radionuclide need to be measured. 27 http://dx.doi.org/10.14311/ap.2016.4.0027 http://ojs.cvut.cz/ojs/index.php/app kamila johnová acta polytechnica ctu proceedings in this method three spectra have to be acquired: the sample itself, the sample with a point source place on its top and the point source without the sample. the method with its limitations is in detail described in [1]. using mathematical tools. with the information technology rapid development the mathematical methods, monte carlo in the first place, will probably soon replace all the other ways of estimating the self-absorption corrections. monte carlo algorithm generally is based on repeated random sampling. in the case of particle transport, the algorithm generates a particle and according to cross section database computes its whole path through the material until it leaves tracked volume or is fully absorbed. with a high number of repetition the final result of this simulation approximates very well the real situation. as the title of this article suggests, in this paper the last item will be further discussed. monte carlo method was previously applied to solve the problem of self-absorption successfully. one of the first works on this topic was [3]. many other articles folowed [4–6], etc. testing different geometries and different monte carlo codes. the following text will describe how this verified method was implemented at our department. 2. experimental equipment many different samples are measured in our laboratory, the most common are: environmental samples (possible contamination monitoring), building materials (radiation protection), soil and rock (geological studies). our equipment includes two hpge detectors designed and produced by canberra company. both of them are classical coaxial germanium detectors with a useful energy range 50 kev to 10 mev. the set up in our laboratory however is suitable for measurement in reduced energy range 50 kev to 3 mev since the laboratory usually works only with natural radioactivity or 137cs contamination. the main part of the detector is ge crystal with ntype and p-type contacts on its surfaces. application of certain potential to the contacts will cause the charge carriers (electrons and holes) move towards the electrodes. due to the very high purity of the ge crystal, only moderate bias is sufficient to “empty” the whole volume between the electrodes, creating so called depleted volume. this region becomes than an active area, where charge carriers, created by photon interactions, are swept by electric field and collected by electrodes. the pre-amplifier, which is incorporated in the cryostat, converts collected charges into voltage pulses. the height of this impulse is proportional to the energy deposited in the volume of the detector by photon interaction. one of the advantages of hpge compared to lithium-drifted detectors (second most common type of semiconductor detector) is that they do not need to be cooled at all time. the detector is stored without cooling, however during the measurement it is cooled in order to prevent thermally generated leakage current. in our laboratory the liquid nitrogen cooling system is used. the contact of ge crystal with the nitrogen (stored in dewar flask) is provided by so called cooling finger made of cu and cover with vacuum. the whole cooling system is called cryostat. the measurement itself can be realized in several geometries of the sample. our laboratory can provide a certified measurement in two geometries. first of them is so called marinelli beaker with a volume 600 ml. this geometry will be denoted as m600 in this article and it is commonly used geometry. the shape of this special container provides very good efficiency of measurement since it surrounds almost whole detector, see figure 1. the samples with volume smaller than 600 ml can be measured using less efficient geometry: a small container (volume 280 ml), which will be denoted as m280 in the rest of this paper. the smaller container can be seen as well in the figure 1. the calibrations of our gamma spectrometry systems are made using certified standards of activity produced by eurostandard cz (czech metrology institute). those standards are the same containers (both m600 and m280) filled with silicon rubber with dispersed known activity of known radionuclide. the spectra are acquired and processed by genietm 2000 analysis software, which is also a product of canberra. 3. monte carlo simulations the monte carlo simulation were implemented using mcnp code. mcnp (monte carlo n-particle) is a general-purpose code developed by los alamos national laboratory. its wide area of application includes radiation protection and dosimetry, shielding against radiation, nuclear criticality safety, detection technique, etc. in discussed problem an mcnp6, the latest release, was used. all geometry plots presented in this article were made using vised – a visual editor for mcnp [2]. 3.1. model of the detector the primer of mc (monte carlo) simulations was model of the detector itself. a rough model was created first using the data provided by the producer of the detector (canberra) in the technical documentation. for more details an x-ray scan of the detector was made. unfortunately some details like thickness of the dead layer or very thin layers generally could not be recognized from the x-ray picture very well. since the model did not still correspond to the reality (calculated and measured efficiencies differed with more than 10 % of relative error), the geometry had to be adapted experimentally. based on the previous experiences the thickness of the dead layer was alternated as was the volume of the vacuum 28 vol. 4/2016 self-absorption corrections based on monte carlo simulations figure 1. 3d visualization of detector model. yellow inner part is the crystal, the top green part represents the m600 (right) and m280 (left) container, blue part is the cooling finger. filled cell. after every change a new set of simulations was made and the results were compared to the measured efficiencies. the agreement of simulated and measured data was checked for m600 and m280 geometry in order to provide the most accurate model of the real detector. the final geometry can be seen in figure 1. 3.2. source and tally the source of radiation in the model was previously described marinelli beaker filled with different materials. both geometries: m600 and m280 were simulated. 3d visualization of the detector with m600 and m280 containers is in figure 2. the simulated photons were generated homogeneously in the whole volume of the container’s filling as it is shown in figure 2. the materials used as a filling of the containers were chosen in order to represent the whole spectrum of samples which are processed in our laboratory. selected materials are: silicon rubber (standards), water, blueberry, wood, soil, rock and brick. the energies of simulated photons were set based on the commonly analysed radionuclides, which are 238u and 232th decay chains, 40k and 137cs. the concentrations of these radionuclides, mainly 238u and 232th, are determined using large number of energy lines originating from different daughter products. all the necessary energy lines were simulated and are part of our database. for the needs of this paper only seven energies were chosen in order to demonstrate the behaviour of correction factor as a function of density and composition. the list of simulated energies is in table 1. in the case of need any other photon energy can be simulated. figure 2. the photons (blue dots) being emitted in the filling of m600 marinelli beaker. choosing tally (i.e. “what should be calculated”) was a simple task, since the objective was to compare simulated data with measured photopeak efficiencies. the best choice for this purpose is f8 type tally, which scores the number of impulses in certain geometry cell. f8 tally (as the other tallies) can be further specified by defining so called energy bins. energy bins provide an information about energy range in which the impulses should be counted. in this case the bin was set so it fully covers the full absorption peak. 29 kamila johnová acta polytechnica ctu proceedings energy [kev] radionuclide decay serie 186.21 226ra 238u 351.92 214pb 238u 661.66 137cs 137cs 727.17 212bi 232th 1460.81 40k 40k 1764.49 214bi 238u 2614.53 208tl 232th table 1. the example photon energies simulated for the needs of this paper. 3.3. data evaluation our approach to the problem allows two different ways of presenting the data. since the previous steps included creating a model of a real detector, complete efficiency curve can be calculated directly for each specific sample. the measured spectra would then be evaluated using this curve. this method is useful for very specific samples or samples that require high precision of measurement. for routine measurements the following way is preferred. the spectra are evaluated using one all-purpose calibration curve estimated by measurement of silicon rubber standards, which were described in the previous text. this process results in the activity a defined for fixed energy e by a = n t · y · ηstandard , where n is number of impulses in the full absorption peak, t is time of measurement, y is the yield of the energy line and ηstandard represents the photo-peak efficiency of the detector for silicon rubber standard. our objective is to estimate correction factor c = c(e,ρ,composition) so the activity of radionuclide in the sample asample can be calculated as asample = a · c = n t · y · ηsample , while ηsample means efficiency of the detector for sample. naturally ηsample = ηsample(e,ρ,composition). it is self-evident that at this point the required correction factor c equals c = ηstandard ηsample . (1) it was previously mentioned, that for fixed geometry the factor c is a function of energy e, density of sample ρ and composition of the material. the importance of those three variables can be forecast using cross section plot. it is one of mcnp option, which allows the user to plot the cross sections for any material and type of particle used in the model. the total cross sections (all interactions included) σt are plotted in figure 3 for all materials discussed in this paper. the detection system in our laboratory can detect photons within the energy range: 50–3000 kev, however energy lines below 100 kev are rarely processed. according to this fact and the information provided in figure 3, it can be assumed that considering the material composition there are only two groups of samples. first group contains “biological materials” or “biomass”, and is represented by wood, blueberries, water and silicon rubber. these materials consists of light elements only (h, o, c) and usually contains a significant amount of water. in can be expected that in the region 100–3000 kev, those samples will not need correction for composition, however the correction for density is still necessary. in another words only one correction curve (the correction factor as a function of density) will be sufficient for all the samples belonging into this group. the other group with similar cross sections contains again natural material but with a certain content of heavier elements (al, si, ca, fe, etc.). in this model this group is represented by brick, soil and rock. again in the region 100–3000 kev only one correction curve will be sufficient for all of these samples. only for lower energies and accurate measurements every material would need a special corrections. 4. results and discussion since this work uses the model of a real detector, it is appropriate at first comment on the comparison of simulated response of the detector with the measured one. it was already mentioned that the model of the detector had been experimentally modified in order to agree to the reality. the comparison of measured and simulated efficiency curves is in figure 4. this plot, made for m600 geometry, indicates a good agreement of measured and simulated data. taking into account the uncertainty of measured data caused mainly by errors in fitting and processing the spectra the model represent very well the real detector. the 5 % uncertainty of measured data is plotted in figure 4 by error bars. similar agreement of measured and simulated data was achieved for m280 geometry. as it was mentioned before the correction factor (1) is a function of composition, material density and energy. in section 3.3 some presumptions about the relation of c with those three parameters were made. in this part these assumptions, based on cross section data, will be confirmed using calculated results. at first the influence of composition of material will be discussed. the cross section plot in figure 3 denoted that the corrections for “biomass” materials as blueberry, wood or any other sample containing mainly h, o and c will be practically the same. the same statement should be valid for group of samples of natural origin containing heavier elements like (al, si, ca, fe, etc.). this group is represented by soil, rock and brick. figure 5 confirms those hypotheses. the correction factors within these two groups were practically the same, with less than 2 % difference. 30 vol. 4/2016 self-absorption corrections based on monte carlo simulations figure 3. the example of cross sections for materials used in the model. figure 4. comparison of measured and simulated efficiency (m600) curve for simulated detector. as a result of this, there are only two curves in this plot, each of them representing one of the groups mentioned above. it is also evident that the differences between the two correction curves are not so significant and for the most of the samples with relatively high uncertainty of measurement (around 10 %) one “mean” correction curve is more than sufficient. the cross section data in the figure 3 also suggested that the photon energy will be affecting the correction factors. the probability of photon interaction (total cross section) was dramatically dropping for low energies (10–70 kev) and continued decreasing slowly for the rest of energy region. this behaviour of cross section for all materials is reflected in figure 6, which plots the correction factor as a function of photon energy. while for the soil-rock types of samples the figure 5. correction factor for m600 as a function of density. composition influence. c is decreasing with energy, for the “biomass” group it is slightly increasing. the dependence of correction factor is logarithmic denoting that for higher energies (above 1500 kev) one correction curve (correction factor as a function of density) will be sufficient. the last parameter which will be commented in this article is a material density. as it was already shown in figure 5 this function is linear. figure 7 presents all correction factors for energies (radionuclides) listed in table 1. these corrections were calculated for m600 filled with brick. however as it was expected based on cross section plot and also proven in the results discussion above, the same curve can be used for soil and rock samples as well. in most of the cases the same corrections can be applied to other samples (wood, water solutions, etc.). 31 kamila johnová acta polytechnica ctu proceedings figure 6. correction factor for m600 as a function of photon energy. figure 7. correction factor for m600 geometry filled with brick. all the previous plots were based on the data calculated for m600 geometry. figure 8 shows the correction factors for smaller m280 geometry as s function of density. even though the behaviour of c = c(ρ) is the same (linear), the corrections itself are higher compared to the m600 geometry. 5. conclusions this article presents a way how the self-absorption problem is dealt with in our gamma spectrometry laboratory. the resulting correction factors c, defined by equation (1), were presented in figure 7, where c was plotted as a function of material density. presented simulations also proved that unless the high precision of measurement is required, the correction factor is independent of material composition. figure 8. correction factor for m280 geometry filled with brick. all the simulations were made using very accurate model of hpge detector. the comparison of measured and simulated efficiencies was plotted in figure 4. this model can be used further more for example for calculations of efficiency for non-standard geometry (geometry which is not calibrated using real certified standard of activity). references [1] g. gilmore. practical gamma-ray spectrometry. 2nd revised edition edition. wiley-blackwell, 2008. [2] r.a. schwarz, et al. graphical user input interface for mcnp. trans. am. nucl. soc. 69:401, 1993. [3] t. nakamura, t. suzuki. monte carlo calculation of peak efficiencies of ge(li) and pure ge detectors to voluminal sources and comparison with environmental radioactivity measurement. nucl. inst. and meth. in phys. res. 205(1):211–218, 1983. doi:10.1016/0167-5087(83)90191-6. [4] o. sima. monte carlo simulation versus semiempirical calculation of autoabsorption factors for semiconductor detector calibration in complex geometries. progress in nuclear energy 24:327–336, 1990. [5] f. sanchez, et. al. a monte carlo based method of including gamma slf-attenuation for the analysis of environmental samples. nuc. ins. and met. in phys. res. 61(4):535–540, 1991. doi:10.1016/0168-583x(91)95334-a. [6] r.m.w. overwater, et al. gamma-ray spectroscopy of voluminous sources corrections for source geometry and self-absorption. nuc. inst. and met. in phys. res. 324(1):209–218, 1993. doi:10.1016/0168-9002(93)90978-q. 32 http://dx.doi.org/10.1016/0167-5087(83)90191-6 http://dx.doi.org/10.1016/0168-583x(91)95334-a http://dx.doi.org/10.1016/0168-9002(93)90978-q acta polytechnica ctu proceedings 4:27–32, 2016 1 introduction 1.1 interactions of gamma radiation with matter 1.2 self-absorption coefficient 2 experimental equipment 3 monte carlo simulations 3.1 model of the detector 3.2 source and tally 3.3 data evaluation 4 results and discussion 5 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2016.3.0015 acta polytechnica ctu proceedings 3:15–18, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app flexibility test of the tools of angioplasty istván hajdu∗, dóra károly, liza pelyhe budapest university of technology and economics, faculty of mechanical engineering, department of materials science and engineering, bertalan lajos u. 7, 1111, budapest, hungary ∗ corresponding author: hajdu999@gmail.com abstract. coronary angioplasty is a procedure used to treat the narrowed coronary arteries. physicians operate with many different tools during the intervention, the main devices are the following: guidewires, guiding catheters, balloon catheters and stents. one of the most important properties of the tools of angioplasty is flexibility. this article introduces a flexibility measuring device and a testing method. with the help of this the flexibility of the tools of angioplasty can be compared easily. keywords: angioplasty, coronary stent, balloon catheter, flexibility. 1. introduction coronary artery disease (cad) is a leading cause of morbidity and mortality in both developing and developed countries [1]. during life fat and cholesterol deposits (called plaques) emerge inevitably in the coronary vessel walls. they gradually increase over time, first they cause coronary stenosis, then complete blockage of the coronary vessels. percutaneous transluminal coronary angioplasty (ptca) is an invasive procedure performed to reduce blockages in coronary arteries [2]. during the process a balloon catheter is guided into the narrowed section of the coronary artery. physicians expand this at the place of the stenosis, thereby the balloon compresses the plaque and ensures continuous flow of blood. usually, a metal mesh is also expanded with the balloon, called stent (fig. 1). stents are commonly made of cocr, ptcr or stainless steel. the most important requirements for stent materials are biocompatibility and hemocompatibility. the stent sholuld not produce a toxic, injurious or immunologic response in living tissue. it is also important for physicians to see the stent during the intervention, therefore it should be visible under x-rays. mri compatibility is another main property of stents, that is why these devices are made from non-magnetic metals (for example austenitic stainless steel). the stent should be flexible to follow the curves of the vessel during the pulsation of the heart, but it should be rigid to compress the plaque [2, 3]. sometimes arteries will still become clogged and narrowed again even with a stent. this phenomenon is called restenosis [4]. low endothelial shear stress can affect the emergence of atherosclerosis and restenosis. stent design, strut thickness and so flexibility have a main role in this field [5]. stent and stent-system properties are mechanical parameters which are important during the deployment, dilation and long term use of the devices. we figure 1. balloon catheter, stent-system and stent. can say that flexibility is one of the most important properties of the tools of angioplasty. the catheter and with it the stent have to follow the curves of the vascular network without damaging or losing their function during deployment, and the stent has to follow the curves of the vessel after dilation at the place of the stenosis. there are many methods to measure the flexibility of the tools of angioplasty, for example the company certiga engineering solutions developed a special equipment. they push the tested tool into a spiral tube as long as it breaks. then they measure the radius at the place of the breaking; they characterize the flexibility with this value [6]. in the study of mori and saito the four-points bending test was utilized. acryl resin bars were attached to both ends of the stent prior to bending measurements, thereby allowing for application of a constant moment to the specimen. through contact with the punch on the acryl resin bars, stent bending with a constant moment and without radial deformation can be achieved [7]. another method is based on the mechanical model of a loaded beam. dr. péter szabadíts characterize the flexibility with the flexural strength of the beam. 15 http://dx.doi.org/10.14311/app.2016.3.0015 http://ojs.cvut.cz/ojs/index.php/app i. hajdu, d. károly, l. pelyhe acta polytechnica ctu proceedings figure 2. the used mechanical model. we also used this model, but by us to the characterization of flexibility the reciprocal of the flexural strength was used [8]. the european standard (en) does not contain any method for the test of the flexibility. the u.s. food and drug administration recommends to wrap the catheter around a series of mandrels with successively smaller radius until the catheter kinks or the lumen collapses. but there is no standard for the stents flexibility measurement [9]. our aim was to design and manufacture a device which is appropriate for the measurement of the flexibility. the method should be fast and easy and adaptable to all kind of tools of angioplasty. 2. materials and methods for the flexibility tests a mechanical model of a loaded beam was used (one end of the cantilever beam was clamped and the other end was loaded; fig. 2). the following equation was used to describe the model: ei = f l3 3f (1) with the help of this equation, the flexibility can be easily calculated: f lexibility = 1 ei (2) for the tests a measuring device was needed (fig. 3), which can be impacted into the testing machine (instron 5965). after designing the device it was fabricated with polyjet and fff rapid prototyping technologies. this device consist of a bottom and a top holder, 33 guide bushes, 2 bolts and 5 spacers. all of the guide bushes had a mortise in their centers with different diameters from 0,4 to 4,7 millimeters. the spacers have the following length: 1, 2, 4, 8 and 16 millimeters; with these between 1 and 31 millimeters all discrete quantities can be assembled. several stents, balloon catheters and stent-systems were investigated as a loaded beam in the model. the investigated devices were the following: 3 balloon catheters with different diameters; 10 stentsystems, from these 6 were the same type, they had difference just in the size of the diameter; and 5 stents figure 3. flexibility measuring device. figure 4. the investigated stents. with different diameters and strut widths (fig. 4). the investigated stents were made from platinum chromium (ptcr) or cobalt chromium (cocr) alloys. the investigations were performed by stereo microscope and testing machine. first, the diameter of the tools was measured with the help of the stereo microscope. then it was impacted into the right size guide bush, which was embed to the top holder. the guide bushes had a bigger outside diameter as the diameter of the mortise on the top holder, so they were fitting tight. after that, the required length was adjusted with the spacers and the whole measuring device with the tested tool was impacted into the testing machine. the tested tools were loaded with a perpendicular force at the end. the load force, the length and deflection of the beam was measured. after the tests the flexibility was calculated with the help of equation (2). every device was loaded five times, between the measurements each was rotated about its axis with the same angles. 16 vol. 3/2016 flexibility test of the tools of angioplasty table 1. the results. figure 5. correlation between flexibility and diameter in the case of stent-systems. 3. results table 1. contains basic information and flexibility of the investigated devices. diameter and length are the nominal sizes. the maximum deflection was 20% of the clamping length, but it was maximized in 3 millimeters. strut width was measured with the help of stereo microscope. three main factors influences the flexibility of the tools of angioplasty, these are the followings: diameter, material and stent pattern. primarily the relationship between the flexibility and diameter was investigated. the results show that between the flexibility and the diameter there is a power function dependence (fig. 5). in the case of stents the relationship between flexibility and strut width was also investigated; between figure 6. correlation between flexibility and strut width in the case of stents. the flexibility and the strut width there is a power function dependence (fig. 6). based on the measurement the flexibility of stents and balloon catheters are about with one order of magnitude greater than the flexibility of stent-systems (fig. 7). 4. conclusions based on the experiments the measuring device is adequate for the investigation of flexibility. in every case the relative standard deviation of the data is lower than 10% (in most cases it is lower than 1%), so the data is homogeneous, this means that it isn’t needed to measure more times in different positions. the flexibility depends on the diameter, raw material, 17 i. hajdu, d. károly, l. pelyhe acta polytechnica ctu proceedings figure 7. correlation between flexibility and diameter of the investigated devices. strut width and stent pattern. the manufacturers don’t give any concrete, numerical data about flexibility of the tools of angioplasty, that’s why these investigations are important. with the help of these results, physicians can easily compare stents or stent-systems according to flexibility, and they can purposefully choose the most suitable tools for the intervention. list of symbols e young’s modulus [mpa] i moment of inertia of beam [mm4] f force [n] l length of beam [mm] f deflection of beam [mm] references [1] k. iwasaki, et al. prevalence of subclinical coronary artery disease in ischemic stroke patients. journal of cardiology 65:71–75, 2015. doi:10.1016/j.jjcc.2014.04.004. [2] g. mani, m. d. feldman, d. patel, c. m. agrawal. coronary stents: a materials perspective. biomaterials 28:1689–1710, 2007. [3] d. jános. az értágítóbetétek anyagainak fejlödése. kohászat 5-6:44–48, 2013. [4] l. petrini, f. migliavacca, f. auricchio, g. dubini. numerical investigation of the intravascular coronary stent flexibility. journal of biomechanics 37:495–501, 2004. [5] k. c. koskinas, et al. role of endothelial shear stress in stent restenosis and thrombosis : pathophysiologic mechanisms and implications for clinical translation. journal of the american college of cardiology 59:1337–1349, 2012. doi:10.1016/j.jacc.2011.10.903. [6] stm-3015 stent system testing ”flexibility” according to din en 14299:2004 7.3.3. http://www.certiga.com/en/stm_3015.en.htm. [7] k. mori, t. saito. effects of stent structure on stent flexibility measurement. annals of biomedical engineering 33:733–742, 2005. [8] p. szabadíts, z. puskás, j. dobránszky. flexibility and trackability of laser cut coronary stent systems. acta bioeng biomech 11:8–11, 2009. [9] class ii special controls guidance document for certain percutaneous transluminal coronary angioplasty (ptca) catheters, 2010 september. http://www.fda.gov/ medicaldevices/deviceregulationandguidance/ guidancedocuments/ucm225145.htm. 18 http://dx.doi.org/10.1016/j.jjcc.2014.04.004 http://dx.doi.org/10.1016/j.jacc.2011.10.903 http://www.certiga.com/en/stm_3015.en.htm http://www.fda.gov/medicaldevices/deviceregulationandguidance/guidancedocuments/ucm225145.htm http://www.fda.gov/medicaldevices/deviceregulationandguidance/guidancedocuments/ucm225145.htm http://www.fda.gov/medicaldevices/deviceregulationandguidance/guidancedocuments/ucm225145.htm acta polytechnica ctu proceedings 3:15–18, 2016 1 introduction 2 materials and methods 3 results 4 conclusions list of symbols references acta polytechnica ctu proceedings doi:10.14311/ap.2016.4.0068 acta polytechnica ctu proceedings 4:68–72, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app measurement of neutron spatial distribution of the bnct epithermal beam at the reactor lvr-15 michaela rabochováa, ∗, miroslav vinšb, jaroslav šoltésb, božena michalcovác a faculty of electrical engineering, czech technical university in prague, czech republic b department of neutron physics, research centre řež, husinec-řež, czech republic c department of neutronography processing, nuclear physics institute of the cas, husinec-řež, czech republic ∗ corresponding author: michaela.rabochova@seznam.cz abstract. in this study, a measurements of neutron field using a special positioning device with a 6li + si detector and image plate is described. the measurements were provided for boron neutron capture therapy (bnct) channel of the lvr-15 reactor in the research centre rez ltd., czech republic. mapping of neutron field represents an essential and crucial part of planning bnct treatment (especially for patients suffering from brain tumor glioblastoma multiforme). keywords: bnct, map of neutron field, 6li + si detector, positioning device. 1. introduction glioblastoma multiforme [1–4] is the most common malignant radio-resistant and chemoresistant tumor of the central nervous system which has been incurable for decades and even today, unfortunately, the prognosis is not favorable [5]. this type of tumor is characterized by very fast growth and aggressive invasion to the surrounding normal brain tissue. it is very difficult to remove all of the affected tumor cells without serious damage of the brain. despite the advanced diagnostics (especially in neuroimaging), and various multimodal therapies that include surgical resection followed by radiation therapy (and alternatively chemotherapy) still remains a considerable problem with finding efficient and gentle treatment of the sensitive area of brain tissue. in this regard, boron neutron capture therapy (bnct) is unique selective radiotherapy based on capture of non-radioactive nuclide 10b by cancer cells and the subsequent capture of thermal neutrons resulting in the nuclear reaction 10b(n, α)7li. these products of reaction selectively damage cancer cells while healthy tissue is spared of radiation load. an absolutely essential part of the treatment planning is determination of parameters and proper setting of correct compensation of neutron beam. the purpose of this project was to measure the spatial distribution of neutrons inside the epithermal horizontal channel of the research reactor lvr-15 (the research centre rez ltd., czech republic) [6], which is used for the method of boron neutron capture therapy. 2. research background 2.1. the special positioning device the mapping of neutron beam was done by using a special positioning device which fixed a si semiconductor detector with 6li converter. the positioning device can moves the detector in three axis and it basically consists of a specially modified support frame with three engines. the special positioning device is shown in fig. 1. figure 1. the special positioning device. 2.2. 6li + si detector for the spatial mapping of neutron beam was used a detector which was consisted of 6li converter and si semiconductor detector [7]. the 6li converter provided a production of 3h (after its insertion into neutron flux) via 6li(n, α)3h reaction. the total energy of the reaction is 4.78 mev (2.73 mev for 3h and 2.05 mev for α particle). the tritium 3h is then detected in a si detector. the active diameter of 6li 68 http://dx.doi.org/10.14311/ap.2016.4.0068 http://ojs.cvut.cz/ojs/index.php/app vol. 4/2016 measurement of neutron spatial distribution of the bnct beam converter is 3 mm; the case of si detector has diameter 25 mm. the distance between converter and detector is approximately 8 mm [8]. the scheme of the detector is shown in fig. 2. figure 2. the scheme of 6li + si detector [8]. 2.3. the image plate another method that has been performed is the experimental determination of the intensity distribution excitations caused by neutron beam by means of the image plates (fig. 3). displaying the neutron flux through the imaging plate is mainly used for neutron radiography method. the principle is based on passing the neutron beam through an object and its subsequent detection on the imagine plate. the result is a 2d (or 3d) image of the different intensities of the neutron flux which is determined by the absorption properties of the object material. in the case of the experiment mentioned in this work was used this method by the same way, but not for displaying a particular subject, but purely for displaying the parameters of the neutron beam intensities. figure 3. the image plate. 3. method of measurement 3.1. the mesaurement with the positioning device the measurement was carried out from march 31 to april 1, 2016 in front of the horizontal channel bnct of the research nuclear reactor lvr-15 in the research centre rez ltd. the acquisition of the data was based on communication between genie-2000 program (detector) and special positioning device utility program. the getting device to work was accomplished by sequential modifications of batch codes. the codes control position of the positioning device and also provide a record of the data of measurements. the main aim of this measurement was to determine the spatial distribution of neutron beam, which is important to verify its homogeneity for subsequent applications of bnct. the map of the route of detector was created and loaded into the control position program of positioning device. the movement of positioning device with 6li + si detector was determined with fixed increments of 1 cm along the horizontal axis y and along the vertical axis z. the time of measurement was ∼ 4 minutes for every position of detector. the scheme of route of the detector is shown in fig. 4. figure 4. the scheme of the route of 6li + si detector. the entire measurement lasted 9 and 12 hours (240 second for one position) and data from detector were analyzed with the software genie-2000 by canberra company [9]. the resulting peak areas of 3h have to be corrected for the change in reactor power as the power was not stable during whole measurement. the power of reactor was higher at the beginning of measurement. for the time of the correction was used the data from ionization chambers which monitoring the bnct channel (see fig. 5). figure 5. the development of reactor power. 69 m. rabochová, m. vinš, j. šoltés, b. michalcová acta polytechnica ctu proceedings the resulting peak areas after correction were calculated based on the following relationship spk = m xi sp, where spk denotes the peak area after correction, m is a median of the numbers of pulses during irradiation, xi is an arithmetic average of the numbers of pulses during the i-th interval of irradiation in a given position and sp is the peak area. 3.2. the measurement with image plate the second experiment of mapping neutron field was performed on april 21, 2016. the main aim was to take a radiography image of thermal neutron beam and its evaluation using neutron imaging plate. the plate was placed under the outlet of the neutron beam. all manipulation were done by hand, the image plate was moved to the beam exit after the beam was opened. this system was provided for pulling the plate upward for start measuring so that the enter of the channel was subsequently overlapped and after the end of irradiation the plate was pull down to the start position. this principle was chosen in order to avoid an unwanted exposure during opening and closing the bnct horizontal channel. start experimental position of image plate is shown in fig. 6. overall exposure time was 2 minutes. figure 6. the neutron image plate (in the red frame) located in the default position below the enter of the bnct horizontal channel (green frame). after the irradiation, the plate was evaluated at the institute of nuclear physics institute of the cas in rez. subsequently, the measured data was processed by the device fujifilm bas-1800 (bas – bioimaging analyzer system), which evaluated the measured intensity. 4. result of measurement 4.1. the mesaurement with the positioning device based on neutron beam radiation of 6li + si detector the map of thermal neutron flux distribution was obtained (fig. 7). values within each field represents the value of the peak area. figure 7. the map of neutron flux distribution. the graphical representation of the measured data is on the fig. 8. the axis x represents the values of the peak areas of 3h and the axes y and z are the positions. the measurement uncertainty was around 1 %. figure 8. 3d graph of neutron flux distribution. 4.2. the measurement with image plate the measured neutrogram of the neutron beam of reactor lvr-15 was then rendered fig. 9. 70 vol. 4/2016 measurement of neutron spatial distribution of the bnct beam figure 9. the neutrongram of neutron beam. the neutron beam is seen here as a red-edge circle. the dark blue color represents the lowest level of excitation (the largest decline), while the white color represents the high intensity. on the neutrongram are partially visible muffled two horizontal bars in the upper and lower half, which are dependent on the properties of a neutron plate and therefore do not reflect the actual intensity values at those locations. unfortunately it is not possible to suppress this phenomenon. small local changes in intensity are the result of small surface scratch of the plate. on the picture is evident symmetrical circular beam (red-purple-white colored intensity) with the lowest attenuation values in the centre, as one might expect. however, the upper half of the image shows a higher level of exposure than the lower half. possible in this case, this could be the influence of gamma radiation to which the neutron plate can be also sensitive. this impact would be affected by the opening bnct channel. another additional measurements intended to clarify these results are planned. it is possible to plot a 3d graph (fig. 10). figure 10. 3d neutrogram of neutron beam. the values on the axes x, y represent the pixel, resp. one unit corresponds to one pixel, which is equivalent to 50 micrometers. thus e.g. 500 px = 2.5 cm, 1000 px = 5 cm, 2000 px = 10 cm etc. the z axis is then plotted excitation intensities in individual pixels. the values are relative and they do not represent actual neutron flux density. the point [x, y] = [0, 0] represents the upper-left corner of the neutron plate placed looking towards the enter of the horizontal channel. again, this graph shows the slope, which is probably caused again by the opening of the channel. 5. conclusions the measurement results show that the profile of the beam is quite homogeneous in its whole crosssection without any significant peaks. in the future, other measurements with other types of converter are planned. also properly functionality of special positioning system which fixed the 6li + si detector was confirmed. the measurement using neutron image plates showed that the neutron beam is symmetrical with the highest intensity in the centre of the beam. it was detected intensity changes in the upper half of the image compared to the lower half of the image. this situation may be caused by the influence of unwanted exposure during opening horizontal channel or by unwanted exposure of gamma radiation at the opening horizontal channel. both methods were sensitive to thermal neutrons (i.e. 6li converter and gd layer of neutron image plate). both experiments confirmed the symmetry of the neutron beam of bnct horizontal channel of lvr-15 reactor. acknowledgements the presented work was financially supported by the ministry of education, youth and sport czech republic project lq1603 (research for susen). references [1] k. urbánek. nádory mozku: gliomy – astrocytom, glioblastoma, příznaky, léčba, diagnostika, 2008. http://www.zbynekmlcoch.cz/informace/medicina/ neurologie-nemoci-vysetreni/nadory-mozkugliomy-astrocytom-glioblastom-priznaky-lecbadiagnostika-prognoza. [2] h. biliková, p. buzrla, j. dvořáčková. problematika mozkových nádorů astrogliální řady. european journal for biomedical informatic 6(1):cs3–cs7, 2010. http://www.ejbi.org/en/ejbi/article/9-csproblematika-mozkovych-nadoru-astroglialnirady.html. [3] m. m. abd-el-barr, e. a. chiocca. how much is enough? the question of extent of resection in glioblastoma multiforme. world neurosurgery 82:e109–e110, 2014. doi:10.1016/j.wneu.2014.05.006. 71 http://www.zbynekmlcoch.cz/informace/medicina/neurologie-nemoci-vysetreni/nadory-mozku-gliomy-astrocytom-glioblastom-priznaky-lecba-diagnostika-prognoza http://www.zbynekmlcoch.cz/informace/medicina/neurologie-nemoci-vysetreni/nadory-mozku-gliomy-astrocytom-glioblastom-priznaky-lecba-diagnostika-prognoza http://www.zbynekmlcoch.cz/informace/medicina/neurologie-nemoci-vysetreni/nadory-mozku-gliomy-astrocytom-glioblastom-priznaky-lecba-diagnostika-prognoza http://www.zbynekmlcoch.cz/informace/medicina/neurologie-nemoci-vysetreni/nadory-mozku-gliomy-astrocytom-glioblastom-priznaky-lecba-diagnostika-prognoza http://www.ejbi.org/en/ejbi/article/9-cs-problematika-mozkovych-nadoru-astroglialni-rady.html http://www.ejbi.org/en/ejbi/article/9-cs-problematika-mozkovych-nadoru-astroglialni-rady.html http://www.ejbi.org/en/ejbi/article/9-cs-problematika-mozkovych-nadoru-astroglialni-rady.html http://dx.doi.org/10.1016/j.wneu.2014.05.006 m. rabochová, m. vinš, j. šoltés, b. michalcová acta polytechnica ctu proceedings [4] a. phuphanich. glioblastoma y astrocitoma maligno. american brain tumor association, 2012. http://www.abta.org/resources/spanish-languagepublications/glioblastoma-y-astrocitomamaligno.pdf. [5] s. s. stylli, a. h. kaye, l. macgregor, et al. photodynamic therapy of high grade glioma – long term survival. j clin neurosci 12(4):389–398, 2005. doi:10.1016/j.jocn.2005.01.006. [6] j. burian. lvr-15 reactor – applications of neutron beam in medicine, biology, dosimetry. iaea technical meeting lecture, vienna, austria, 2011. [7] l. viererbl, j. šoltés, m. vinš, et al. measurement of thermal neutron beam parameters in the lvr-15 research reactor. in transaction of igorr 2013 conference. 2013. http://www.igorr.com/home/ liblocal/docs/igorr2013/07_1001.pdf. [8] l. viererbl, j. burian, s. hladky, et al. si diode with converter used for measurement of epithermal neutron beam of lvr-15 reactor. nuclear instruments and methods in physics research section a: accelerators, spectrometers, detectors and associated equipment 580(1):366–368, 2007. doi:10.1016/j.nima.2007.05.180. [9] canberra industries, inc. genie 2000 basic spectroscopy software, 2016. http://www.canberra. com/products/radiochemistry_lab/pdf/g2kbasicspect-ss-c40220.pdf. 72 http://www.abta.org/resources/spanish-language-publications/glioblastoma-y-astrocitoma-maligno.pdf http://www.abta.org/resources/spanish-language-publications/glioblastoma-y-astrocitoma-maligno.pdf http://www.abta.org/resources/spanish-language-publications/glioblastoma-y-astrocitoma-maligno.pdf http://dx.doi.org/10.1016/j.jocn.2005.01.006 http://www.igorr.com/home/liblocal/docs/igorr2013/07_1001.pdf http://www.igorr.com/home/liblocal/docs/igorr2013/07_1001.pdf http://dx.doi.org/10.1016/j.nima.2007.05.180 http://www.canberra.com/products/radiochemistry_lab/pdf/g2k-basicspect-ss-c40220.pdf http://www.canberra.com/products/radiochemistry_lab/pdf/g2k-basicspect-ss-c40220.pdf http://www.canberra.com/products/radiochemistry_lab/pdf/g2k-basicspect-ss-c40220.pdf acta polytechnica ctu proceedings 4:68–72, 2016 1 introduction 2 research background 2.1 the special positioning device 2.2 6li + si detector 2.3 the image plate 3 method of measurement 3.1 the mesaurement with the positioning device 3.2 the measurement with image plate 4 result of measurement 4.1 the mesaurement with the positioning device 4.2 the measurement with image plate 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0047 acta polytechnica ctu proceedings 5:47–50, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app current challenges for research activities in the field of railway infrastructure otto plasek brno university of technology, faculty of civil engineering, veveri 95, brno, czech republic correspondence: plasek.o@fce.vutbr.cz abstract. the increasing importance of railway systems, arising from both national and european strategic documents, leads to increasing demands on its infrastructure. transport development in the european territory is defined in the strategy summarized in the white paper, which understands the rail sector as a key sector in terms of sustainability of the transport development due to the suppression of congestions, carbon oxide emissions, reliability and safety of transport. the paper is aimed in the specific technical aspects of rail infrasructure, which emphasize their ability to contribute to an efficient and sustainable transport in europe and justify motivations for increasing the attractiveness of rail transport. current challenges for research activities in the field of railway infrastructure that will led to achievement of the objectives which were defined in the white paper and further specified in the particular requirements of rstakeholders in the railway industry are discussed. keywords: railway infrastructure, railway research, interoperability of railway infrastructure, noise and vibrations. 1. introduction the increasing importance of railway systems, arising from both national and european strategic documents, leads to increasing demands on its infrastructure. transport development in the european territory is defined in the strategy summarized in the "white paper roadmap to a single european transport area towards a competitive and resource efficient transport system", which was released by the european commission in 2011 [1]. this strategy defines the "roadmap to a single european transport area" with the aim of creating a competitive and efficient transport system. it is evident that an efficient transport systems are crucial for the competiveness of european enterprises in the global economy. this fact is supported by an argument that transport and freight storage costs are 10-15 % of the final products costs. every european household gives approximately 13.2 % of its budget on transport-related products and services. annual congestion costs in europe are approximately 1 % of gross domestic product. transport development strategy is based on several obvious assumptions. it is expected a continuous increasing of the oil price which should be at least twice in 2050 compared to 2005, while transport is dependent on oil for at least 96 %. greenhouse gas emission should be reduced 80 % bellow the level of 1990 to limit the global temperature change to 2c. alarming is also fact that exceeding of capacity limits for road and air transport requires annually 1 % additional costs in europe. the increase of freight transport up to 40 % in 2030 and up to 80% in 2050 is assumed. moreover, the infrastructure is developed unevenly in the eastern and western parts of european union. nowadays only about 4,800 km of motorways, no high speed railway lines and conventional railway lines often in poor conditions are in the new member states [2]. the white paper for transport understands the rail sector as a key sector in terms of sustainability of the transport development due to the suppression of traffic congestions, carbon oxide emissions, reliability and safety of transport. the necessity of transformation of rail transport to become more attractive and to increase its market share of passenger and freight transport for medium distance (up to 300 km) in 2050 is defined in this strategic plan. in this respect, the white paper puts a very ambitious goal of moving the road transport to rail or water transport in volume of 50 % [1]. the specific technical aspects of rail transport, which emphasize their ability to contribute to an efficient and sustainable transport in europe and justify motivations for increasing the attractiveness of rail transport are stated bellow in the paper. 2. current role of rail transport rail passenger transport in europe now plays an important role in long-distance transport (from 300 km) through especially high-speed trains [3]. the definition of high-speed services is not easy and involves a wide range factors which determine operator access to the passenger rail transport. nevertheless, it is usually considered the rail traffic at speeds higher than 230 km/h. the rail passenger services not quite successfully compete in individual transport for medium 47 http://dx.doi.org/10.14311/app.2016.5.0047 http://ojs.cvut.cz/ojs/index.php/app otto plasek acta polytechnica ctu proceedings distances up to 300 km so far, because the quality of service reflects considerable time losses when access to rail transport and while waiting for the connection or transfer. the exception is train connection in directions where there is no motorway available. that is why rail operators is really interested in the connection praha ostrava. it should be noted that in the mid-range distances will play a crucial role, yet unmentioned aspects of the definition of high-speed traffic, which express the quality of transport: • infrastructure (including civil works, tracks, overhead catenary systems, etc.) • comfortable railway stations (their location, functional design, equipment, etc.) • rolling stock (technically advanced, comfortable, modern design etc.) • operation (planning, management, regulations) • advanced command and control systems (etcs) • sophisticated strategy of maintenance • coherent financing, marketing and management the suburban transport in big cities or regional centers is another important area of the positive benefits of passenger rail transport. the inclusion of railway transport into integrated transport systems reflects this trend. rail freight transport currently plays an important role in the field of long distance, especially the transcontinental transport. the effectiveness of the railways in this regard improves the use of combined systems. however, it must be noted that the increase in the medium-range transport (mainly domestic) will always run into tough competition with road transport. 3. research activities in the field of railway infrastructure research, development and innovation will be focused on construction of high-speed lines, increasing speeds on existing rail lines, increasing reliability and durability of tracks with mixed railway traffic in the context of increasing service load and capacity. the current issue of interoperability of the railway infrastructure is taken into particular consideration. rail infrastructure is in the state that corresponds a history of its construction, maintenance, renovation and modernization. rail infrastructure must be considered as a very complex system, which consists of components that are of different technical, technological advancement and age. railway track consists both from earthworks such as embankments or cuts and from advanced railway superstructure both ballasted and ballastless, characterized by the controlled vertical stiffness of substructure and with resilient rail fastening. most of tracks are of different age and technological advancement due to the gradual replacement of the structural elements except completely renewed or modernized tracks. the main activities of research activities aimed in the development of rail infrastructure, both railway and urban lines are focused on requirements of speed increase, higher service and axle load, safety and security, ride comfort while meeting economic needs, environmental requirements, i.e. reduction of noise, suppression of vibrations spreading, especially in urban areas, also reducing costs of maintenance and repair works, reducing energy consumption and consumption of raw materials. calls for research projects both national and european correspond to the above specified requirements. the focus of the research activity in the czech republic is the preparation of technological solutions for the design and construction of high-speed railways. an important part of the process is development of decision-making strategies for choosing the most appropriate design solutions for high speed lines (eg. ballasted vs. ballastless track). research activities are focused on the preparation of new structural and technological solutions to meet the requirements of the railway superstructure and substructure of high-speed lines, especially on the structural design of switches and crossings. the attention must be paid to the particular design of the high speed structures as well as to encourage the readiness of domestic suppliers in railway industry and track work contractors. it is also important to pay attention to the management and maintenance systems of high speed lines, focusing on the question both of logistics and policy-warranty service and preventive maintenance in defined cycles. regarding railway substructure, earthworks were constructed usually in the 19th century. structural layers ensuring the strength of substructure corresponding to the increased axle load were constructed only for certain lines in the second half of the 20th century. at the turn of the millennium corridor lines have been modernized, where structural layers in the active zone of the embankment have been constructed or completely renewed. most sections outside the corridor lines does not comply with current requirements due to the age of the earthworks and their original design load. remedying this situation will require high costs and be associated with a number of construction works, which require service disturbance or interruption. most of the research is therefore concentrated on finding technologies, increasing strength of substructure especially for weak, compressible or unstable soils. the solutions are searched with help static and dynamic analyses and an optimization of the stiffness of the railway substructure, development of advanced mobile diagnostic methods and tools. solutions which help to minimize service disturbances and interruptions at the least possible cost of works and consumption of materials are looked for. railway superstructure was many times renewed in comparison with rail substructure in the same section. whenever rail superstructure was usually renewed with up to date technology. this means that except 48 vol. 5/2016 current challenges for research activities in the field of railway infrastructure figure 1. test section in plana nad luznici railway station under sleeper pads corridor lines railway superstructure is a mixture of components of different age and technology along the track. it can be state that currently used technology of ballasted tracks are comparable with ones used in high speed lines. design of switches and crossings (s&c) is the only exception. however, the domestic s&c manufacturer in its applied research activities intensively develops new technologies. an optimized geometry, the application of the elastic elements in order to suppress the influence of variable vertical stiffness along the length of the switch and crossing, development of resistant materials for crossings, operational systems of the new generation, trough sleepers, monitoring systems for switches and movable frogs are the subject of research activities. a certain barrier to the development is the fact that high speed switches crossings is not easy tested by high-speed traffic in the czech republic. rail infrastructure administration faces with the development of rail defects which are usually connected to the increase of axle load and service speed. this can be remedied by increasing the rail steel quality, but in the surface layer resisting to fatigue loading. thus new types of rail defects occur and appropriate preventive measures are looked for. the short pitch corrugation occurs in lower rail in tight curves which during wheelset passages besides vibrations causes also noise emission to the track vicinity. the only effective measure is the grinding or milling of the rails, which is very costly, requires service interruption and significantly reduces the life span of the rails. effective measures may consist in an additional elasticity of the track or modification of the friction coefficient at the wheel-rail contact. a significant part of the research work is aimed at the technology procedures and management of construction activities and maintenance work. next subject of research works in this field is the development of advanced technological processes, rules of guarantee service and preventive maintenance strategy. innovative and advanced technological procedures for figure 2. investigation of rail corrugation by the salamander device havlickuv brod maintenance and reconstruction of tracks, logistics issues, management and strategy of maintenance are developed. 4. environmental aspects of rail infrastructure outcomes of research activities lead to the development of such infrastructure design, which also contributes to the reduction of negative impacts on the track vicinity, while solutions are being sought especially economically efficient from a lifecycle. cost perspective. railway tracks mainly in urban areas are perceived as a significant source of noise and vibrations. with a growing number of rail vehicles is adversely affected track vicinity. this negative influences were worse if any measures would not be used. the negative effects in some cases led to an idea to shift of tracks away of city centre, which, however, caused more difficult access of passengers to rail transport. this approach has been modified but this requires significant suppression of noise emission and vibrations propagation. the noise caused by traffic not only adversely contributes to the emergence of various diseases. in principle, the noise generated by the rail transport can originate from wheel-rail contact, from vehicle engines and aggregates, which is significant for low service speed, and aerodynamic noise, which is significant for high service speed. the noise originated from wheelrail contact rolling, squeezing, impact is a common trouble for every rail vehicle which increases necessity to deal with this phenomenon. for this reason, not only an influence of parameters of rolling stock but also railway superstructure and substructure parameters and service conditions are investigated. vibrations that spread to building structures and building foundations are other loads. vibrations are induced by dynamic effects from the train movement across the track irregularities. characteristics of dynamic effects are related to number, weight and tech49 otto plasek acta polytechnica ctu proceedings figure 3. installation of noise barriers in usti nad orlici nical state of vehicles, service speed, running dynamic, track structure, track quality, track alignment, geological parameters, etc. preventive measures at the vibration source are most economical and most effective. if it is not possible to prevent structural vibrations at a source, then this source should be completely separated from adjacent structures. special elastic rail fastening, under sleeper pads, under ballast mats, mass-spring systems or ballastless track are used as measures in track structure for vibration mitigation. most types of these measures are available today as deliveries of recent applied research projects. higher initial costs, which are usually outweighed by non-financial benefits, are currently an obstacle for immediate application of measures preventing vibration propagation. 5. conclusions railway infrastructure administration put emphasize on research and development and implementation of research results into practice. it is considered as the fundamental way to increase the competitiveness of rail transport in the process of developing a sustainable and efficient transport. stakeholders (manufacturing companies, contractors, design companies, universities and research centres), which participate in the investment production and maintenance of the railway infrastructure in the czech republic, in the areas of infrastructure, energy, control, command and signalling, have jointed their efforts in the czech technology platform. its objective is the linkage of the scientific and technical potential of universities, research and project institutes together with the production potential of construction and manufacturing companies for the implementation of the following areas of activities [4]: • support of innovation and increase of competitiveness of the members of the association • structuring and support of implementation of development, research and testing projects ensuring the existing production of the members of the association with the requirements of the technical specifications of interoperability of the trans-european railway system in the sub-systems of infrastructure, energy, control, command and signalling • acquiring of financial funds for implementation of these projects • international (european) activities related to the creation of new regulations for construction, production and maintenance and related tests and evaluation of the railway industry a variety of projects supported by national funding agencies, particularly the technology agency of the czech republic, grant agency of the czech republic or by the ministry of industry and trade are solved nowadays. the project "centre for efficient and sustainable transport infrastructure -cesti" belongs among the most important projects, aimed in railway infrastructure. the ministry of education, youth and sports promotes the development of science and research projects through operational programs, an example is the project "interoperability of railway infrastructure competence network (iricon)". projects can be also supported from different european funding programmes. the leading european companies in railway industry joined efforts in a joint undertaking shift2rail within the eu framework programme for research and innovation ? horizon2020. the shift2rail has just opened the first calls for project of applied research for its members and in the form of opencalls to non-members, too. with respect to the current support of science and research projects at national and european level, with regard to available facilities of newly built research centres, excellent research teams, it can be assumed that research activities will achieve the objectives which were defined in the white paper and further specified in the particular requirements of railway companies, contractors and manufacturers in the railway industry. acknowledgements the article was processed under financial support of the project lo1408 "admas up advanced building materials, structures and technologies" supported by ministry of education, youth and sports within the "national programme for sustainability i". references [1] european commission. white paper on transport. towards a competitive and resource efficient transport system., 2011. http://www.transforum-project.eu/ transforum/white-paper-on-transport.html. [2] union internationale chemins de fer (uic). rail and sustainable development, 2001. http://www.uic.org. [3] union internationale chemins de fer (uic). high speed rail, fast track to sustainable mobility, 2009. http://www.uic.org. [4] www.sizi.cz, 2016. http://www.sizi.cz/introduction. 50 http://www.transforum-project.eu/transforum/white-paper-on-transport.html http://www.transforum-project.eu/transforum/white-paper-on-transport.html http://www.uic.org http://www.uic.org http://www.sizi.cz/introduction acta polytechnica ctu proceedings 5:47–50, 2016 1 introduction 2 current role of rail transport 3 research activities in the field of railway infrastructure 4 environmental aspects of rail infrastructure 5 conclusions acknowledgements references 38 acta polytechnica ctu proceedings 1(1): 38–41, 2014 38 doi: 10.14311/app.2014.01.0038 fullerenes, pahs, amino acids and high energy astrophysics susana iglesias-groth1,2 1istituto de astrofisica de canarias, la laguna, via láctea sn, 38201 la laguna, spain 2departamento de astrof́ısica de universidad de la laguna, la laguna, tenerife, spain corresponding author: sigroth@iac.es abstract we present theoretical, observational and laboratory work on the spectral properties of fullerenes and hydrogenated fullerenes. fullerenes in its various forms (individual, endohedral, hydrogenated, etc.) can contribute to the uv bump in the extinction curves measured in many lines of sight of the galaxy. they can also produce a large number of absorption features in the optical and near infrared which could be associated with diffuse interstellar bands. we summarise recent laboratory work on the spectral characterisation of fullerenes and hydrogenated fullerenes (for a range of temperatures). the recent detection of mid-ir bands of fullerenes in various astrophysical environments (planetary nebulae, reflection nebulae) provide additional evidence for a link between fullerene families and diffuse interstellar bands. we describe recent observational work on near ir bands of c60 + in a protoplanetary nebula which support fullerene formation during the post-agb phase. we also report on the survival of fullerenes to irradiation by high energy particles and gamma photons and laboratory work to explore the chemical reactions that take place when fullerenes are exposed to this radiations in the presence of water, ammonia and other molecules as a potential path to form amino acids. keywords: ism molecules: fullerenes, pahs nebulae. 1 introduction the discovery of fullerenes in 1985 when trying to reproduce the chemistry of the atmospheres of red giant stars (kroto et al., 1985) led to the identification of the third allotropic form of carbon. the most abundant fullerene molecule produced in those experiments was c60, a hollow molecule with 60 carbon atoms distributed in 12 pentagons and 20 hexagons (truncated icosahedron symmetry). comparetively, we recall polycyclic aromatic hydrocarbons (pahs) are planar molecules consisting of carbon rings and hydrogen, these rings are similar to benzene. the naphthalene and anthracene are the most simple pahs with two and three benzene rings, respectively. these molecules have been postulated as potential carriers of the unidentified infrared emission bands and of the diffuse interstellar bands which are ubiquitous in the interstellar medium (see e.g. léger and puget 1984). similarly, fullerenes and their hydrogenated forms, are potential carriers of the interstellar bands. several studies have shown that carbon ring based molecular forms are very stable against uv radiation. this is remarkable because the basic structures of amino acids and in general of the molecular basis of life are essentially conformed by such carbon rings making them rather stable against possible mutations. laboratory work bychen et al. (2008) have aaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaa aaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaa demonstrated that under strong uv radiation a mixture of naphthalene, ice and methane at low temperature, it is possible to form a large variety of amino acids. fullerenes with carbon atom number of 20 (m2 + n2 + nm) where n and m are integers have icosahedral symmetry groups ih and exhibit high stability. these molecules are very stable against uv/gamma radiation and collisions with other particles. fullerenes have been detected in carbonaceou chondrites, a type of meteorite that originated in the early solar system,with abundances of the order of 0.1 ppm (pizzarello et al. 2000). on earth, fullerenes have been detected in sedimentary layers of the cretaceous-tertiary boundary (ktb) in china and bulgaria and in the mineral shunghite of the region of karelia (russia). in the various phases of the interstellar medium it is likely the presence of both, fullerenes and hydrogenated fullerenes, the socalled fulleranes (cnhm). a review of the properties of fullerenes and fulleranes can be found in cataldo & iglesias-groth (2009, 2010). here, recent developments on the astrophysical search for ionised fullerenes and on the properties of fullerenes when exposed to high energy radiation. 38 http://dx.doi.org/10.14311/app.2014.01.0038 fullerenes, pahs, amino acids and high energy astrophysics 2 astrophysical searches for fullerenes theoretical work on the photoabsorption spectrum of icosahedral fullerenes suggests that these molecules can be responsible of the strongest feature in the interstellar extinction curve, the uv bump at 2175 å. for large fullerenes and buckyonions the experimental photoabsorption spectrum is very poorly known and the potential role of these molecules in interstellar absorption has been explored on a theoretical basis (see, for example, iglesias groth et al., 2002, 2003). in these works a theoretical approach to the photoabsorption specturm was based on a hückel and pariser-parr-pople (ppp) model. fits of the computed spectra to measurements of the interstellar extinction in the uv/optical for different lines of sight in our galaxy (fitzpatrick, 1999) were obtained (see details of the model can be found in iglesias-groth, 2004). very good fits to the extinction curves are provided by the photoabsorption cross sections computed for both individual fullerenes and buckyonions. the comparison of these models with the observed uv bump leads to an estimate of the number density of these molecules (in the range 0.2-0.08 ppm) and of the percentage of carbon locked in fullerenes and buckyonions in the ism (iglesias-groth 2004, 2007) which depending the size distribution of these molecules could reach up to 25 % of all carbon in the ism. fullerenes c60 and c70 present active vibrational bands in the near infrared (3-20 µm) which have been key for their identification in planetary nebulae, reflection nebulae and post-agb stars (see e.g. cami et al. 2010, zhang and kwok, 2011). experimental measurement of the molar absorptivity for these infrared transitions is essential for the determination of fullerene abundances in such astrophysical contexts. laboratory work is being conducted to measure the dependence of the molar absorptivity of fullerene infrared bands with temperature from approx. 100 k to 500 k to cover a variety of conditions in the interstellar medium (iglesias-groth et al. 2011). 2.1 new results on ionised fullerenes: c+60 recently, it was reported the detection of mid-ir vibrational transitions of the fullerene c60 in a carbon rich protoplanetry nebula, iras 01005+7910 (zhang and kwok 2010). we recorded the spectrum of this object between 5500 and 10000 å with resolving power of 57,000 and detected two bands at 9577 and 9632 å which are consistent with laboratory measurements of the c60 cation (see figure 1. iglesias-groth and esposito, 2013). if these two bands were produced by c+60 in the material surrounding the central post-agb star we could infer that ∼ 1 % of carbon is trapped in this ionized form of fullerenes and that the cation abundance is much higher than that of the neutral fullerene species. it appears that ionized fullerene species in this protoplanetary nebula are significantly more abundant than in neutral form, which is consistent with the predictions of ionization models of pahs (bakes and tielens 1995, salama et al. 1996) for irradiated clouds near a hot star. figure 1: the two near ir bands in the spectrum of iras01005+7910, attributed to the cation of c60 (iglesias-groth & esposito 2013) these observations provide additional evidence for the presence of fullerenes in protoplanetary nebulae and suggest that a significant production takes place in this late stage of stellar evolution. mid-ir bands of c+60 could be present in the 7-20 µm spectrum of iras 01005+7910 and are also likely to be detected in the spectra of planetary nebulae. accounting for the ionized species of fullerenes appears necessary in nebulae with a nearby source of ionizing photons. a caveat: the derived abundances and excitation temperatures of planetary nebulae with claimed detections of fullerenes may have to be revised if indeed there is a significant contribution of cations to the formation of the bands used in these analyses. high resolution spectroscopy will be required for a reliable determination of the relative abundance of neutral and ionized fullerenes in these objects. 2.2 the stability of c60 and c70 towards copuscular and γ radiation the stability of c60 and c70 fullerenes in the interstellar medium and embedded in meteorites and comets has been investigated with γ irradiation and with he+ ion bombardment. the radiation dose generated by radionuclides decay expected to be delivered to fullerenes buried at a depth of ≥ 20 m in comets and meteorites is about 3 mgy per million years. we have exposed fullerenes to various radiation doses and found that these molecules are by far resistant to this type of radiation. laboratory measurements on the concentra39 susana iglesias-groth tion of samples of fullerenes exposed to strong radiation doses indicate that these molecules can survive millions of years inside comets and meteorites (cataldo et al. 2009). this provides a natural explanation to their presence inside certain carbonaceous chondrites. in this laboratory experiments, fullerenes were also exposed to intense particle bombardment. fullerens adsorbed or deposited on the surface of carbon are exposed to cosmic ray bombardment with estimated doses of radiation comprised between 30 and 65 mgy per 106 years. we carried out experiments to test the survival of both c60 and c70 fullerenes and found that the complete amorphization occurs at about a radiation dose of 250 mgy. thus we infer that after 4-8 millions of years exposure to cosmic rays it is expected to producce a complete amorphization (cataldo et al. 2009). 2.3 gamma radiolysis of c60 fullerene in water and water/ammonia mixtures; relevance of fullerenes fate in ices of interstellar medium in order to explore the potential astrochemical role of fullerenes,iglesias-groth et al. (2013) have explored the radiolysis of fullerene c60 dispersed in h2o, h2o/nh3, h2o/methanol and h2o/nh3/methanol under gamma radiation doses of 250 and 500 kgy. it was found that c60 originally insoluble in the above mentioned hosting matrix became soluble as a consequence of multiple hydroxylation and oxidation reaction produced by the free radicals generated by the radiolysis of the hosting matrix. the changes undergone by c60 were studied by infrared spectroscopy (ft-ir) and by electronic absorption spectroscopy. the astrochemical consequences of the present study are that c60 ejected in the interstellar medium for instance from protoplanetary and planetary nebulae can condense together with water and other ices in dense molecular clouds. under the action of high energy radiation c60 reacts with the free radicals generated from the matrix where it is embedded it is solubilized and consequently its carbon content becomes available for further abiotic processes of synthesis of molecules of astrobiological interest. the behavior of c60 appears comparable to that of common pahs which are also hydroxylated and oxidized under similar conditions. 3 stability toward high energy radiation of non-proteinogenic amino acids: implications for the origins of life in order to investigate the response of amino acids to high energy radiation, cataldo et al. (2013) have taken a series of non-proteinogenic amino acids, most of them found frequently in carbonaceous chondrites, and exposed them to solid state radiolysis in vacuum to a total radiation dose of 3.2 mgy. this corresponds to 23% of the total dose expected to be taken by organic molecules buried in asteroids and meteorites since the beginning of the solar system in 4.6 x 109 years. the radiolyzed amino acids where studied by ft-ir spectroscopy, differential scanning calorimetry (dsc) and by polarimety and optical rotatory dispersion (ord). it is found that an important fraction of each type of amino acid is able to ”survive” to the massive dose of radiation and also the enantiomeric excess is partially preserved. based on these results it is concluded that it is not a surprise to find amino acids even in enantiomeric excess in carbonaceous chondrites. 4 conclusions the astrochemical consequences of the recent discoveries of fullerenes in various astrophysical contexts are not fully understood yet. fullerenes are far resistant to high doses of gamma radiation and bombardment by high energy particles. it is likely that fullerenes ejected in the interstellar medium,for instance from protoplanetary and planetary nebulae can condense together with water and other ices in dense molecular clouds. under the action of high energy radiation fullerenes react with the free radicals generated from the matrix where they are embedded. if c60 is trapped in water ices it is hydroxylated and oxidized by the radiolysis products of water. the oxidation of c60 makes this molecule hydrophilic and hence soluble in water. the same phenomenon occurs in water/ammonia, in water/methanol and in water/ammonia/methanol mixtures. thus, c60 which in the solid state displays a considerable radiation resistance, when embedded in radiolytic sensitive matrices like those just mentioned, it reacts swiftly, it is solubilized and consequently its carbon content becomes available for abiotic processes of synthesis of other molecules of astrobiological interest. this compares well with the behavior of pahs which are also hydroxylated and oxidized by the free radicals produced by the radiolysis of the hosting matrix (gudipati & allamandola 2006, ricca et al. 2007, ashbourn et al. 2007, cuylle et al. 2012, nuevo et al. 2012). it is known that pahs and ices exposed to strong uv radiation can lead to the formation of amino acids. fullerenes exposed to gamma radiation provides a path for a variety of astrochemical reactions which have to be explored in much more detail, which ultimately could lead to the formation of amino acids, which are likely to survive under the intense gamma radiation. further laboratory work is essential to explore the role of high energy radiation in astrophysical sources and carbon molecules as a route to prebiotic molecules. 40 fullerenes, pahs, amino acids and high energy astrophysics references [1] bakes, e. & tielens, a. g. g. m. 1995, aspc, 73, 59 [2] cami, j, bernard-salas, j, peeters, e & malek, s. e. 2010, science [3] cataldo, f., iglesias-groth, s. 2009, mnras, 400, 291 doi:10.1111/j.1365-2966.2009.15457.x [4] cataldo, strazzulla and iglesias-groth et al. 2009, mnras, 394, 615 [5] cataldo, f & iglesias-groth,s 2010, in fullerenes:the hydrogenated fullerenes, springer, berlin [6] chen et al. , 2008, mnras, 384, 60 [7] cuylle, s. h., tenenbaum, e. d., bouwman, j., linnartz, h. & allamandola, l. j. 2012, mnras, 423, 1825 doi:10.1111/j.1365-2966.2012.21006.x [8] fitzpatrick, e. l 1999, pasp, 111, 63 [9] iglesias-groth, s. et al., 2002, jchph, 116,1064810655 [10] iglesias-groth, s. et al., 2003, jchph, 118, 71037111 [11] iglesias-groth, s. 2004, ,apjl 608, l37 doi:10.1086/422216 [12] iglesias-groth, s. 2007, apjl, 661,l167 doi:10.1086/518832 [13] iglesias-groth, s., cataldo, f., ursini, o.& manchado, a. 2011, mnras, 210, 1447-1453. [14] iglesias-groth,s.,cataldo, f. & manchado, a. 2011 , mnras, 413, 213 doi:10.1111/j.1365-2966.2011.18124.x [15] iglesias-groth, s. & esposito, m. 2013,apjl ,776, l2 doi:10.1088/2041-8205/776/1/l2 [16] iglesias-groth, s., hafez, y., angelini, g. & cataldo, f. 2013, in press [17] kroto, h. w. et al. 1985, nature, 318, 162 [18] léger, a. & puget, j. 1984, a&a, 137,5 [19] nuevo, m., milam, s. n.& sandford, s. a. 2012, asbio 12, 295 [20] pizzarello, s. & cronin, j. r. 2000,gecoa,64, 329 doi:10.1016/s0016-7037(99)00280-x [21] ricca, a., bakes, e. l. o. & bauschlicher, c. w., jr. 2007, apj, 659, 858 doi:10.1086/512037 [22] salama, f., bakes, e. l. o., allamandola, l. j.& tielens, a. g. g. m. 1996, apj, 458, 621 [23] zhang, y., kwok, s. 2011,apj , 730, 126 doi:10.1088/0004-637x/730/2/126 41 http://dx.doi.org/10.1111/j.1365-2966.2009.15457.x http://dx.doi.org/10.1111/j.1365-2966.2012.21006.x http://dx.doi.org/10.1086/422216 http://dx.doi.org/10.1086/518832 http://dx.doi.org/10.1111/j.1365-2966.2011.18124.x http://dx.doi.org/10.1088/2041-8205/776/1/l2 http://dx.doi.org/10.1016/s0016-7037(99)00280-x http://dx.doi.org/10.1086/512037 http://dx.doi.org/10.1088/0004-637x/730/2/126 introduction astrophysical searches for fullerenes new results on ionised fullerenes: c60+ the stability of c60 and c70 towards copuscular and radiation gamma radiolysis of c60 fullerene in water and water/ammonia mixtures; relevance of fullerenes fate in ices of interstellar medium stability toward high energy radiation of non-proteinogenic amino acids: implications for the origins of life conclusions 200 acta polytechnica ctu proceedings 1(1): 200–204, 2014 200 doi: 10.14311/app.2014.01.0200 high-resolution x-ray spectroscopy of galactic supernova remnants satoru katsuda1, hiroshi tsunemi2 1riken (the institute of physical and chemical research), 2-1 hirosawa, wako, saitama 351-0198, japan 2department of earth and space science, graduate school of science, osaka university, 1-1 machikaneyama, toyonaka, osaka, 560-0043, japan corresponding author: katsuda@crab.riken.jp abstract high-resolution x-ray spectroscopy of galactic supernova remnants (snrs), based on grating spectrometers onboard xmm-newton and chandra, has been revealing a variety of new astrophysical phenomena. broadened oxygen lines for a northwestern compact knot in sn 1006 clearly show a high oxygen temperature of ∼300 kev. the high temperature together with a lower electron temperature (kte ∼ 1 kev) can be reasonably interpreted as temperature non-equilibration between electrons and oxygen behind a collisionless shock. an ejecta knot in the puppis a snr shows blueshifted line emission by ∼1500 km s−1. the line widths are fairly narrow in contrast to the sn 1006’s knot; an upper limit of 0.9 ev is obtained for o viii lyα, which translates to an oxygen temperature of kto < 30 kev. the low temperature suggests that the knot was heated by a reverse shock whose velocity is ∼4 times slower than that of a forward shock. anomalous intensity ratios in o vii heα lines, i.e., a stronger forbidden line than a resonance line, is found in a cloud-shock interaction region in puppis a. the line ratio can be best explained by the charge-exchange emission that should arise at interfaces between the cold/warm clouds and the hot plasma. there are several other targets for which we plan to analyze high-quality grating data prior to the operation of the soft x-ray spectrometer onboard astro-h. keywords: supernova remnants x-rays individuals: sn 1006, puppis a. 1 introduction there are many scientific motivations to perform highresolution x-ray spectroscopy of supernova remnants (snrs), since the x-ray emission is often dominated by thin thermal emission, namely line emission. for instance, accurate line widths allow us to measure the degree of temperature equilibration between electrons and ions as well as efficiency of particle acceleration. resolved fine structures of lines give us insight into emission processes. we can also reveal ejecta dynamics and measure abundances in unprecedented detail. current challenges are based on grating spectrometers onboard xmm-newton and chandra. both of the grating systems are slitless. therefore, while they work well for point-like sources or compact extended sources, they are generally not suitable for largely extended sources because off-axis emission along the dispersion direction is detected at wavelength positions shifted with respect to the on-axis source. most snrs in the large/small magellanic clouds are small enough for the gratings, and indeed over 10 papers have been published in literature (e.g., burrows et al. 2000; rasmussen et al. 2001; behar et al. 2001; van der heyden et al. 2001; 2002; 2003; flanagan et al. 2004; kosenko et al. 2008; 2011; broersen et al. 2011). a problem here is that we can not deduce spatial information, since the grating spectra are integrated for the entire snr. this causes serious ambiguities of interpretations. therefore, it is important to observe large galactic snrs for spatiallyresolved high-resolution spectroscopy. however, in this case, grating spectra suffer from spectral degradation due to the spatial extent of the source. the only solution to this dilemma is to focus on locally bright and compact features in large snrs. we here review a few successful examples of high-resolution spectroscopy of galactic snrs with the xmm-newton’s reflection grating spectrometer (rgs: den herder et al. 2001). the rgs has a large dispersion angle and is more suitable for extended sources than chandra’s grating spectrometer, although there are some nice results (lazendic et al. 2006; rutherford et al. 2013). 2 fast-moving knots in sn 1006 and puppis a a pioneering work employing the rgs to obtain highresolution spectra from galactic snrs was given by vink et al. (2003) who observed a northwestern (nw) knot in sn 1006 (see fig. 1 (a)). as shown in fig. 1 200 http://dx.doi.org/10.14311/app.2014.01.0200 high-resolution x-ray spectroscopy of galactic supernova remnants (b), the rgs spectrum showed several emission lines, including o vii heα forbidden and resonance, from the knot. by taking account of the degradation effect due to spatial extent, vink et al. (2003) found intrinsic broadening of o k-shell lines to be σ = 3.4 ± 0.5 ev. this value was later revised by the same group based on additional deep observations to σ = 2.4±0.3 ev (broersen et al. 2013). the amount of broadening corresponds to an oxygen temperature of 275+72−63 kev, if attributed to thermal doppler broadening. on the other hand, the electron temperature is derived to be ∼1.35 kev, which is ∼200 times lower than the oxygen temperature. this temperature discrepancy is the evidence for temperature non-equilibration as is expected behind collisionless shocks; the ion temperatures (kti) immediately behind collisionless shocks are given by kti =3/16miv 2 sh, where mi is an elemental mass and vsh is a shock speed, so that the initial temperatures are proportional to particle masses. 18 19 20 21 22 23 0 0.5 1 c ou nt s s− 1 å − 1 wavelength (å) knot knot bek (a) sn 1006 (c) puppis a (b) (d) o viii lyα r f figure 1: (a) x-ray view of sn 1006 in 500–599 ev, taken from broersen et al. (2013). the rgs field of view is within the two lines. (b) rgs spectra zooming into o vii heα from the nw knot in sn 1006. the positions of the resonance (21.6å), the intercombination (21.8å), and the forbidden line (22.1å) are indicated. this figure is taken from broersen et al. (2013). (c) x-ray view of the puppis a snr in 0.5–5 kev. the rgs spectra taken from the two solid lines (hosting the knot) or dashed lines (hosting the bek) are shown in fig. 1 (d) and fig. 3, respectively. (d) rgs spectrum zooming into o k-shell lines from the ejecta knot in puppis a. in the model, the contribution from the knot is indicated as a blue line. rest-frame positions of the o viii lyα (18.9å) and the o vii heα lines are indicated as dashed lines. the blueshifts are clearly visible. along the same line, another interesting target is an ejecta knot in the northeast of the puppis a snr (see, fig. 1 (c)) whose southern portion is positionally coincident with an optical o-rich filament (the so-called ω filament which is not shown here: winkler & kirshner 1985). based on nondispersive x-ray ccds onboard xmm-newton (epic: turner et al. 2001, strüder et al. 2001), katsuda et al. (2008) noticed blueshifted k-shell lines from the knot. however, the poor energy resolution of the epic data (e/∆e ∼ 20) prevented them from conclusive arguments; the doppler velocity ranged from 1700+700−800 km s −1 at the south of the knot, which agrees with the optical measurement of ∼1500 km s−1 (winkler & kirshner 1985), to 3400+1000−800 km s −1 at the north of the knot. since the knot is bright and compact (∼5 times brighter than its surroundings and the size is smaller than ∼3′), the rgs is capable of yielding a highresolution spectrum at a level of e/∆e ∼ 150. thus, we performed a new xmm-newton observation, and ob201 satoru katsuda, hiroshi tsunemi tained a nice rgs spectrum as shown in fig. 1 (d). it exhibits prominent k-shell lines, including o vii heα and o viii lyα. the lines are clearly blueshifted by 1480±140±60 km s−1 (the first and second term errors are measurement and calibration uncertainties, respectively), which is fully consistent with optical measurements for the ω filament. it is also found that doppler velocities are uniform within the knot (katsuda et al. 2013b). in addition, the rgs spectra enabled us to measure line broadening to be σ < 0.9 ev at o viii lyα, indicating an upper limit of an oxygen temperature of 30 kev. interestingly, the temperature for puppis a’s knot is at least an order-of-magnitude lower than that for sn 1006’s knot, even though they are comparably fast moving. to interpret the dramatic difference in oxygen temperatures between sn 1006’s knot and puppis a’s knot, we investigate temperature equilibration through coulomb interactions behind collisionless shocks. in the calculation, we assume elemental abundances and the ionization timescales given in the literature (table 1 in broersen et al. 2013 for sn 1006 and table 2’s case-b in katsuda et al. 2013b for puppis a). also, since the degree of electron heating at collisionless shocks is not yet understood well, we examined various initial electronto-proton temperature ratios, te/tp = β, for each shock speed. in this way, we ran calculations for several shock speeds. figure 2 shows the best-representative models, where temperature curves are illustrated only for electrons (black) and oxygen (red). for comparison, we also show model curves with β = me/mp. we find that the data require fairly different shock speeds; a fast (∼3000 km s−1) shock for sn 1006 and a relatively slow (∼600 km s−1) shock for puppis a. also, the values of β are obtained to be ∼0.06 for sn 1006 and less than ∼0.5 for puppis a. these values roughly agree with an observational trend shown in van adelsberg et al. (2008). furthermore, the value for sn 1006 is consistent with a direct measurement from hα spectroscopy for the optical filament (β < 0.07: ghavamian et al. 2002). on the other hand, it is interesting to note that our best-estimated value of 0.06 is higher than a recent theoretical prediction, β = 0.023 for vsh > 1500 km s −1 (ohira & takahara 2008). we will revisit this issue in our future work. the inferred shock velocity for sn 1006 is roughy consistent with the proper motion of the shock ahead of the knot (winkler et al. 2003; katsuda et al. 2013a). therefore, a simple interpretation is that the knot was heated by a forward shock. however, we cannot rule out a possibility of reverse-shock heating. in fact, the knot shows an elevated si abundance, suggesting that it originates from the sn ejecta that should have been heated by a reverse shock. in this case, the (reverse) shock velocity should be higher than ∼5000 km s−1 (see, the discussion in katsuda et al. 2013b). such a fast shock requires an oxygen temperature of ∼800 kev. this is significantly higher than the x-ray measurement, thereby a significant fraction of the shock energy may go into cosmic-rays. if true, this suggests efficient cosmic-ray acceleration at the reverse shock. as for puppis a, the inferred shock velocity, ∼600 km s−1, is less than the shocked-gas velocity, ∼2000 km s−1, derived by the doppler velocity of ∼1500 km s−1 and the optical proper motion of ∼1250 km s−1 at a distance of 2.2 kpc (winkler et al. 1988). this is consistent with a picture that the ejecta knot was heated by a reverse shock rather than a forward shock, since the forward shock velocity should be ∼2700 km s−1 (= 4/3 × 2000 km s−1 according to the rankine-hugoniot relation). 109 1010 1011 1012 1 10 100 kt ( ke v ) net (cm −3s) oxygen electro n vshock = 3000 km s −1 solid line: β = 0.06 dotted line: β = me/mp 109 1010 1011 1012 0.1 1 10 kt ( ke v ) net (cm −3s) oxygen electro n vshock = 600 km s −1 solid line: β = 0.5 dotted line: β = me/mp sn 1006 puppis a figure 2: temperature histories for electrons (black) and oxygen (red) as a function of net. left and right panels are for the nw knot in sn 1006 and the ejecta knot in puppis a, respectively. 202 high-resolution x-ray spectroscopy of galactic supernova remnants 3 cloud-shock interaction regions in puppis a possible sites for high-resolution spectroscopy of galactic snrs are not only ejecta knots but also swept-up ism regions. in particular, cloud-shock interaction regions in the puppis a snr are promising, because they are bright and compact (e.g., hwang et al. 2005). so far, rgs spectra taken from the bright eastern knot (bek) and the northern knot were presented by katsuda et al. (2012). in fig. 3, we show one of the bek spectra focusing on the o vii heα lines, which was taken from the region within the dashed lines in fig. 1 (c). interestingly, we see a strong forbidden line with respect to the resonance line. this is inconsistent with any thermal emission models with reasonable plasma parameters (i.e., kte = 0.3–0.7 kev and net = 1010–5×1011 cm−3 s). after considering a few possibilities to reproduce the anomalous line ratio, katsuda et al. (2012) proposed that the charge-exchange emission, that arises at interfaces between the cold/warm clouds and the hot plasma, can best explain the data. this would support a long-standing expectation of chargeexchange x-ray emission in snrs. 21 21.5 22 22.5 0 0.5 1 c ou nt s s− 1 å − 1 wavelength (å) puppis a bek f i r figure 3: rgs spectrum zooming into o vii heα from the bek in puppis a. the positions of the resonance (21.6å), the intercombination (21.8å), and the forbidden line (22.1å) are indicated. 4 future prospects there are several nice rgs data sets of galactic snrs that are waiting for being analyzed. these include bright shells of rcw 86, an equatorial belt in g292.0+1.8, and a southwestern knot in the cygnus loop. in addition, it is worth while to search for other promising targets and observe them. we are planning to analyze the rgs data in the near future. the rgsbased study is a great pathfinder for the coming highresolution x-ray spectroscopy of snrs with the nondispersive soft x-ray spectrometer (mitsuda et al. 2010) onboard astro-h (takahashi et al. 2012). acknowledgement s.k. is supported by the special postdoctoral researchers program in riken. this work is partly supported by a grant-in-aid for scientific research by the ministry of education, culture, sports, science and technology (23-000004 and 24-8344). references [1] behar, e., et al. 2001, a&a, 365, l242 [2] broersen, s., et al. 2011, a&a, 535, 11 [3] broersen, s., et al. 2013, a&a, 552, 9 [4] burrows, d. n., et al. 2000, apj, 543, l149 doi:10.1086/317271 [5] den herder, j. w., et al. 2001, a&a, 365, l7 [6] flanagan, k. a., et al. 2004, apj, 605, 230 doi:10.1086/382145 [7] ghavamian, p., et al. 2002, apj, 572, 888 doi:10.1086/340437 [8] hwang, u., et al. 2005, apj, 635, 355 doi:10.1086/497298 [9] katsuda, s., et al. 2008, apj, 678, 297 doi:10.1086/586891 [10] katsuda, s., et al. 2012, apj, 756, 49 doi:10.1088/0004-637x/756/1/49 [11] katsuda, s., et al. 2013a, apj, 763, 85 doi:10.1088/0004-637x/763/2/85 [12] katsuda, s., et al. 2013b, apj, 768, 182 doi:10.1088/0004-637x/768/2/182 [13] kosenko, d., et al. 2008, a&a, 490, 223 [14] kosenko, d., et al. 2011, a&a, 532, 114 [15] lazendic, j. s., et al. 2006, apj, 651, 250 doi:10.1086/507481 [16] mitsuda, k., et al. 2010, spie, 7732, 773211 [17] ohira, y., & takahara, f. 2008, apj, 688, 320 doi:10.1086/592182 [18] rasmussen, a. p., et al. 2001, a&a, 365, l231 [19] rutherford, j., et al. 2013, apj, 769, 64 doi:10.1088/0004-637x/769/1/64 203 http://dx.doi.org/10.1086/317271 http://dx.doi.org/10.1086/382145 http://dx.doi.org/10.1086/340437 http://dx.doi.org/10.1086/497298 http://dx.doi.org/10.1086/586891 http://dx.doi.org/10.1088/0004-637x/756/1/49 http://dx.doi.org/10.1088/0004-637x/763/2/85 http://dx.doi.org/10.1088/0004-637x/768/2/182 http://dx.doi.org/10.1086/507481 http://dx.doi.org/10.1086/592182 http://dx.doi.org/10.1088/0004-637x/769/1/64 satoru katsuda, hiroshi tsunemi [20] strüder, l., et al. 2001, a&a, 365, l18 [21] takahashi, t., et al. 2012, spie, 8443, 84431z [22] turner, m. j. l., et al. 2001, a&a, 365, l27 [23] van adelsberg, m., et al. 2008, apj, 689, 1089 doi:10.1086/592680 [24] van der heyden, k. j., et al. 2001,a&a,365,l254 [25] van der heyden, k. j., et al. 2002,a&a,392,955 [26] van der heyden, k. j., et al. 2003,a&a,406,141 [27] vink, j., et al. 2003, apj, 587, l31 doi:10.1086/375125 [28] winkler,p.f. & kirshner,r.p. 1985,apj,299,981 [29] winkler, p. f., et al. 1988, iau colloq. 101: 65 [30] winkler, p. f., et al. 2003, apj, 585, 324 doi:10.1086/345985 204 http://dx.doi.org/10.1086/592680 http://dx.doi.org/10.1086/375125 http://dx.doi.org/10.1086/345985 introduction fast-moving knots in sn 1006 and puppis a cloud-shock interaction regions in puppis a future prospects acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0017 acta polytechnica ctu proceedings 5:17–21, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app design parameters of buffer stops petr guziur brno university of technology, faculty of civil engineering, veveří 331/95, 602 00 brno correspondence: guziur.p@fce.vutbr.cz abstract. paper discuses reasons of building buffer stops and situations that may occur in railway station leading to build safety tracks. also discusses parameters of buffer stops that enter its design, such as collision speed and kinetic energy absorbing capacity. furthermore, presents categories of buffer stops depending on principles of absorbing the kinetic energy and points pros and cons of each structure. keywords: buffer stop, kinetic energy, braking force, dead-end track. 1. introduction buffer stop is a device at the end of dead-end track or closed track with a purpose to stop the rolling stock. in czech republic it is allowed to use three types of buffer stops (according to national regulation čd ž9 [1]). all three types are fixed (rigid) construction, yet there are numerous designs of buffer stops used abroad. we can divide buffer stops in categories depending on the principles of absorbing the kinetic energy, which is the most important parameter of buffer stops. as basic categories we can name fixed buffer stops, hydraulic buffer stops and friction buffer stops. every type has its pros and cons and suitable place to be installed (depending on circumstances). 1.1. flank protection one of the places, where buffer stop has to be installed, is the end of safety track which represents direct flank protection of train paths. there are two types of flank protection of train path: • direct flank protection – protection of rolling stock from non-permitted ride of rolling stock from side approaching track – flank protection switch or derail is used, • indirect flank protection – protection of rolling stock from non-permitted ride of rolling stock from side approaching track – signals with prohibiting signal is used. and there are different conditions for flank protection of train path with speed up to 120 km.h−1 and above this speed. direct flank protection of train path with speed higher 120 km.h−1 is required: • from all sidings, • from all service tracks. in this case, rolling stock arriving from siding or service track should head to the safety track and probably has low speed at its end (by the buffer stop). another situation with possibility of higher impact speeds may occur on safety tracks from running tracks (see chapter 1.1.1). 1.1.1. mutually excluded train path amongst other conditions, safety and signalling plant of second or higher category has to forbid the simultaneous setting of train path for which the train path with higher speed than 120 km.h−1 meets, crosses or overlaps with continuation of another train path. this condition is crucial for direct flank protection. if there is no direct flank protection of the train path with speed higher than 120 km.h−1, one of train paths is not allowed to be set or the speed is restricted up to 120 km.h−1. speed restriction depends on possibilities of main signals (what speed can be signalled). in railway station, where restricted speed is used, the restriction brings speed to lower level, than in railway station, where no restriction is needed (with speed of 120 km.h−1). this issue causes regular delay of trains on daily bases, since not even system like kango is able to cope with this situation. the only result is to build flank protection switches not only from sidings and service tracks but even from running tracks where threat of the train path with speed higher than 120 km.h−1 occurs. 2. buffer stop’s design parameters as a crucial parameter for designing buffer stop is its kinetic energy absorbing capacity, thus stopping a rolling stock of certain mass with a certain collision speed. there are various situations of buffer stop’s location, therefore it is needed to approach the design individually. whether the buffer stop is at the end of a dead-end track in station by platform, or ends a safety track (either from sidings, service track or even running tracks). each situation gives us different speed and different type of train (its mass) that can run into a buffer stop. as mentioned above, amount of kinetic energy that buffer stop is able to absorb is the crucial parameter for its design. kinetic energy of moving rolling stock can be calculated as sum of the kinetic energy of transitional motion and kinetic energy of rotating parts of the rolling stock (1.). 17 http://dx.doi.org/10.14311/app.2016.5.0017 http://ojs.cvut.cz/ojs/index.php/app petr guziur acta polytechnica ctu proceedings vehicle groups vehycle types ρ[-] trains regular passengers or freight trains 0,06electric motor unit 0,15 0,20 cars cars with mechanical traction transition 0,12 0,15 cars with traction motors 0,20 0,025 passengers cars 0,04 0,06 full freight cars 0,04 0,05 empty freight cars 0,1 0,12 locomotives steam 0,08 0,10 electric 0,20 0,30 motor 0,15 0,30 table 1. coefficient of rotating parts [2] speed [km.h−1] speed [m.s−1] impact mass [t]100 200 300 400 500 600 700 800 900 1000 10 2,78 386 773 1159 1546 1932 2319 2705 3091 3478 3864 15 4,17 869 1739 2608 3478 4347 5217 6086 6956 7825 8694 table 2. examples of kinetic energy [kj] (1.) ekin,c = ekin,p + ekin,r (2.) ekin,p = 1 2 mv2 (3.) ekin,r = 1 2 iω2 with substitution equation (2.) and (3.) into equation (1.) and following modification it is possible to calculate kinetic energy of moving rolling stock, using equation (4.). (4.) ekin,c = 1 2 m(1 + ρ)v2 • ekin,c ... total kinetic energy [j]; • ekin,p ... kinetic energy of transitional motion [j]; • ekin,r ... kinetic energy of rotating parts [j]; • m ... weight of rolling stock [kg]; • v ... collision speed [m.s−1]; • i ... moment if inertia [kg.m2]; • ω ... angular speed [s−1]; • ρ ... coefficient of rotating parts [-]. examples of kinetic energy are given in table 2 for buffer stop, as a device (structure) that has to work properly and must ensure high level of safety and reliability, we need to involve a safety coefficient in calculations. therefore, the buffer stop’s kinetic energy absorbing capacity has to be determined as follows: r >= ekin,ck • r ... buffer stop’s kinetic energy absorbing capacity [j]; • ekin,c ... total kinetic energy [j]; • k ... safety coefficient [-]. in table 3 are shown examples of safety coefficient graded by the level of protection and type of train according to austrian standards [3]. 2.1. collision speed the collision speed is defined as the maximum permissible speed in which trains may travel when colliding with buffer stop. there are various approaches, how to determine collision speed. german standard ds 800 01 [4] use the collision speed based on train type: • passengers trains: 15 km.h−1 • freight trains: 10 km.h−1 austrian standard dv b 53 [3] use the collision speed based on track type: • main line trains: 15 km.h−1 • empty passengers trains or shunting: 10 km.h−1 take an example how the collision speed can be calculated. driver is responsible to make a proper braking according to signals. by error he overlooks a presignal and has no information about decelerating the train. he starts to decelerate at the point of signal with prohibiting signal using the emergency brake. table 4 displays braking distances (using operating brakes and emergency brake) and speeds of trains at the end of dead-end track (model situation: assuming the distance between signal with prohibiting signal and buffer stop is 100 m). parameters for table 4: • * train will stop before the buffer stop, • os deceleration – 0,50 m.s−2 (2x unit 451/452) [5], • r, ex, ic deceleration – 0,45 m.s−2 (locomotive 363 + 8 cars type y) [5], 18 vol. 5/2016 design parameters of buffer stops type of train and security level k [-] passengers trains 1,5 freight trains and shunting 1,2 freight trains and shunting, when it is necessary to protect various systems which are located behand or nearby buffer stop 1,5 freight trains and shunting, in cases where there are traffic zones, structures or residential houses located behind or nearby buffer stop 1,8 preventing the fall of any train or rolling stock into abyss 2,0 table 3. safety coefficient [3] speed [km.h−1] braking distance [m] speed at the end of dead-end track [km.h−1] os r, ex, ic emergency braking os r, ex, ic emergency brake 50 192,90 214,33 40,19 35 37 * 60 277,78 308,64 57,87 48 49 * 70 378,09 420,10 78,77 60 61 * 80 493,83 548,70 102,88 71 72 13 90 625,00 694,44 130,21 82 83 43 100 771 857,34 160,75 93 94 61 table 4. deceleration of trains buffer stop type 6 cars of 15 t 1 car of 80 t rail buffer stop 1,0 km.h−1 1,6 km.h−1 concrete type "sudop" 0,7 km.h−1 1,1 km.h−1 concrete type "dsb" 1,0 km.h−1 1,6 km.h−1 table 5. resistances of buffer stops used in czech republic [1] • emergency brake deceleration – 2,40 m.s−2. 3. types of buffer stops 3.1. fixed buffer stops with mechanical bumpers fixed (rigid) buffer stop is one of the most used types of buffer stops considering its history as an oldest one. construction of fixed buffer stop consists of a block or frame fixed rigidly to the rails or the ground. as other constructions, fixed buffer stop has its pros and cons. one of advantages is that it can be placed at the end of the dead-end track, thus it does not reduce the usable length of the track. however, in this case cons prevail, such as low resistance and manner of deceleration. resistances (as a state of usability) of fixed buffer stops used in czech republic are shown in table 5. if the mass or speed is higher, either the buffer stop or the train is destroyed. assuming both buffer stop and rolling stock frame will be reinforced significantly, and therefore cannot deform, the deceleration figure 1. example of hydraulic buffer stop (oleo comp.) [7] is unacceptable (see table 6). 3.2. hydraulic buffer stop (fixed buffer stop with hydraulic bumpers) hydraulic buffer stops are similar in construction to fixed. consist of a block or frame fixed rigidly to the rails or the ground. hydraulic buffer stops absorb kinetic energy in gradual manner (depends on the type of the hydraulic bumpers). if the energy is higher than the bumper is able to absorb, buffer stop is destroyed. parameters of the buffer stop on figure 1: • kinetic energy absorbing capacity: 2688 kj, • bumpers stroke: 2400 mm. 3.3. friction buffer stop friction buffer stop is the most effective way to stop moving rolling stock. the way of absorbing the kinetic energy, thus decelerating the rolling stock is the most efficient and safe. not only the train decelerates in 19 petr guziur acta polytechnica ctu proceedings impact speed [km.h−1] impact speed [m.s−1] stroke of mechanical bumpers [mm] [6] 75 105 decelartion [m.s−2] overload decelartion [m.s−2] overload 5 1,39 6,43 0,66g 4,59 0,47g 10 2,78 25,72 2,62g 18,37 1,87g 15 4,17 57,87 5,90g 41,34 4,21g table 6. deceleration of train and overload on passengers on impact the fixed buffer stop figure 2. length needed for buffer stop placement and to ensure braking distance [8] gradual manner over a longer distance (time period), but the kinetic energy absorbing capacity could be very high. friction buffer stop generally consist of rigid steel frame with buffers, connected to rails using arresting devices (friction jaws). in case of collision, kinetic energy is transformed into heat by means of friction. therefore, energy absorbing capacity of friction buffer stop depends on the number of friction jaws, friction coefficient resp. braking force of jaws and length of braking. nevertheless, there is a disadvantage. friction buffer stop needs a braking distance, therefore cannot be placed at the end of the deadend track and shortens its usable length. moreover the track behind the buffer stop needs to be horizontally straight and contain no welds and joints or other obstacles over rails. notes for figure 2: • lv ... length needed for buffer stop placement and to ensure braking distance; • lb ... buffer stop length; • lw ... maximum braking distance. three types of friction buffer stops can be named: • friction buffer stop (without additional brakes), see figure 3; • friction buffer stop with additional brakes, see figure 4; • friction buffer stop with hydraulic bumpers (with/without additional brakes). notes for figure 3: (1.) collision triangle; (2.) breaking devices; (3.) buffers; (4.) device for returning the buffer stop to working position after collision; (5.) reinforcement; figure 3. friction buffer stop without additional brakes [8] figure 4. friction buffer stop with additional brakes [8] (6.) reinforcement. notes for figure 4: (1.) additional arresting devices for added braking; (2.) steel profile below rails for reinforcement of the track; (3.) jointed connection belts between arresting devices; (4.) lateral connection between arresting devices. 3.3.1. braking force braking force is the determining factor of friction buffer stop. as mentioned above, buffer stop must absorb high amount of kinetic energy. calculation of breaking force in case of friction buffer stops depends on numbers of arresting devices, braking force of each arresting device and maximum braking distance. r = nbfblw • r ... buffer stop’s kinetic energy absorbing capacity [j]; 20 vol. 5/2016 design parameters of buffer stops figure 5. rawie 16 zeb, stopping distance/impact speed/impactmass diagram, safety coefficient 1,0 and 1,5 [9] • nb ... number of arresting devices [-]; • fb ... braking force of a single arresting device [n]; • lw ... maximum braking distance [m]. braking force of friction buffer stop with additional brakes comes from formula up there. formula is modified considering braking force of additional brakes while braking distance of each additional arresting device is included separately. r = nz∑ i=1 2fbilw i • r ... buffer stop’s kinetic energy absorbing capacity [j]; • nz ... number of a pair of additional arresting devices [-]; • fbi ... braking force of a single arresting device, based on a braking distance [n]; • lw i ... length of braking distance of a pair of arresting devices [m]. an example of friction buffer stop’s braking effectiveness is shown in figure 5 (friction buffer stop without additional brakes with absorbing capacity of 640 kj.m−1). 4. conclusion many aspects have to be taken in account while designing buffer stop. one has to consider its location – ending of dead-end track in railway station, ending safety track from sidings, service track or running track. every scenario brings different requirements such as type of train, collision speed, impact mass etc. in general, three types of buffer stops are used, depending of the principle of absorbing the energy. fixed buffer stops are more than useless for higher collision speeds considering its manner of deceleration and its resistances. however, using fixed buffer stops is justified e.g. in shunting yards, where are low speeds and no passengers on board. as more appropriate ending of dead-end track is usage of hydraulic or friction buffer stop. resistances of those types are much higher than fixed buffer stops. both this construction decelerates train in gradual manner and no harm to the rolling stock or buffer stop itself is done if the design was precise. acknowledgements the paper was created with support of the project no. lo1408 "admas up advanced materials, structures and technologies" supported by the ministry of education, youth and sports within the targeted support of program "national program for sustainability i". references [1] čd ž9 železniční spodek, vzorový list železničního spodku, zarážedla. české dráhy, s.o., divize dopravní cesty, o. z. praha, 2001. in effect from: 2002-04-01. [2] j. široký. mechanika v dopravě i âăş kolejová vozidla [online]. updated 2003, [2016-01-13], http://homen.vsb.cz/~s1i95/mvd/skr_mvd.pdf. [3] ’́obb: dv b 53. die gestaltung von oberbauanlagen. [4] db: ds 800 01. bahnanlagen entwerfen allgemeine entwurfsrichtlinien. [5] l. fiala. provozní dopady aplikace ochranných vzdáleností podle tnž 34 2620 master thesis. university of pardubice, pardubice. supervisor pavel drda, 2010. [6] tnž 28 2605 kolejová vozidla – železniční. trubkové nárazníky s korýtkovým vedením. typy, základní parametry, technické požadavky, zkoušení. nymburk: čsd, 1991. in effect from: 1991-07-01. [7] oleo end stops. oleo international [online]. updated 2015-07-09, [2015-11-07], http://www.oleo.co.uk/products/end-stops. [8] israel railways ltd. railway buffer stops planning guidlines [online], 2009. updated 2013, [201511-10], http://www.iroads.co.il/sites/default/ files/imce/ir_buffer_stops_guidelines.doc. [9] rewie bahntechnik strassenbahn. rawie gmbh & co. kg [online]. updated 2015-08-20, [2015-11-07], http://www.rawie.de/index.php/de/bahntechnik/ strassenbahn. 21 http://homen.vsb.cz/~s1i95/mvd/skr_mvd.pdf http://www.oleo.co.uk/products/end-stops http://www.iroads.co.il/sites/default/files/imce/ir_buffer_stops_guidelines.doc http://www.iroads.co.il/sites/default/files/imce/ir_buffer_stops_guidelines.doc http://www.rawie.de/index.php/de/bahntechnik/strassenbahn http://www.rawie.de/index.php/de/bahntechnik/strassenbahn acta polytechnica ctu proceedings 5:17–21, 2016 1 introduction 1.1 flank protection 1.1.1 mutually excluded train path 2 buffer stop's design parameters 2.1 collision speed 3 types of buffer stops 3.1 fixed buffer stops with mechanical bumpers 3.2 hydraulic buffer stop (fixed buffer stop with hydraulic bumpers) 3.3 friction buffer stop 3.3.1 braking force 4 conclusion acknowledgements references 212 acta polytechnica ctu proceedings 2(1): 212–216, 2015 212 doi: 10.14311/app.2015.02.0212 detection of diatomic molecules in the dust forming nova v2676 oph m. nagashima1, a. arai1,2, t. kajikawa1, h. kawakita1, e. kitao1, t. arasaki1, g. taguchi1, y. ikeda1 1kyoto sangyo university, koyama astronomical observatory 2university of hyogo, nishi-harima astronomical observatory corresponding author: g0837714@gmail.com abstract novae are generally considered to be hot astronomical objects and show effective temperatures up to 10,000 k or higher at their visual maximum. but, it is theoretically predicted that the outer envelope of the nova outflow can become cool enough to form molecules that would be dissociated at high temperatures. we detected strong absorption bands of c2 and cn radicals in the optical spectrum of nova v2676 oph, a very slow nova with dust formation. this is the first report of the detection of c2 and the second one of cn in novae during outburst. although such simple molecules are predicted to form in the envelope of the outflow based on previous studies, there are few reports of their detection. in the case of v2676 oph, the presence of the molecular envelope is considered to be very transient, lasting several days only. keywords: cataclysmic variables classical novae optical spectroscopy ir individual: v2676 oph. 1 introduction dust formation in the outflow of a nova had been proposed by mclaughlin (1935) to explain the rapid drop in the visual light-curve of dq her in 1934. dust formation in fh ser had also been confirmed by infrared observations by geisel et al. (1970). dq her was the first nova in which molecular absorption bands of cn in optical wavelength had been identified. the formation of molecules as the precursor to dust grains in novae is considered important for understanding how dust grains form in the outflow of novae. in the case of dq her in 1934, strong cn absorption bands of both violet and red systems had been detected merely 2 days after the visual maximum, and these absorption bands were identified for only 1 week approximately (wilson & merrill, 1935; sanford, 1935; stoy & wyse 1935; antipova 1969; sneden & lambert, 1975). since the formation of simple molecules such as cn is considered an intermediate process in the formation of dust grains from the hot atomic gas in the outflow of a nova, molecular formation in the early phase of dq her might be associated with dust formation its later phase. although cn is the first molecule observed in a nova during outburst (in the case of dq her), there are no further reports about cn in other novae. whereas, carbon monoxide (co) emission in the early phase of novae has been observed by both photometric and spectroscopic observations. in particular, the first overtone band of co (∆v = 2) has been routinely detected in near-infrared spectra of several novae, as reviewed by evans & rawlings (2008). based on previous observations, the correlation between detection of co and dust formation is noticeable. however, hydrogen (h2) or other molecules have not detected in novae during the early phase of their outburst. here we report the detection of c2 and cn in optical spectra of the classical nova v2676 oph during the early phase of its outburst. this nova could be classified as a slow nova, and it showed a rapid drop in its visual light curves about 90 days after its discovery. this is the first report of the detection of c2 in novae during outburst and the second for cn. furthermore, co emission in the near-infrared had been detected in this nova (rudy et al., 2012). 2 observations nova v2676 oph (pnv j17260708-2551454) was discovered at ut 2012 mar 25.789 (t = 0 day) by h. nishimura (reported in the central bureau electronic telegram (cbet) 3072). after the discovery of the nova, we carried out spectroscopic observations with the low-dispersion spectrograph losa/f2 (shinnaka et al., 2013) mounted on the 1.3-m araki telescope at koyama astronomical observatory on ut 2012 mar 27. on the 212 http://dx.doi.org/10.14311/app.2015.02.0212 detection of diatomic molecules in the dust forming nova v2676 oph first night, we detected narrow balmer emission lines (both hα and hβ) and narrow fe ii emission lines on a highly reddened continuum that seemed to be due to interstellar extinction. the color excess e(b − v ) was estimated by using the balmer decrement and the color of the nova (b−v ); it was determined as 0.71±0.02 and 0.72±0.06, respectively. based on the p cygni profile of the hα emission, the expansion velocity was estimated to be ∼ 800 km/s. we concluded that the object was a fe ii-type classical nova in the early phase (arai & isogai, 2012). after the first observation of v2676 oph, we continued the observation of it routinely to assess its the spectroscopic evolution (nagahsima et al., 2014). 4500 5000 5500 6000 6500 7000 7500 n o rm a liz e d s p e ct ru m wavelength [å] ∆v=+1 ∆v=0 ∆v=−1 c 2 ∆v=+5 ∆v=+4 ∆v=+3 cn hβ na i hα o i v2676 oph tx psc figure 1: comparison of the spectrum of v2676 oph obtained on apr 8 with that of a typical carbon star, tx psc. tick marks indicate telluric absorption lines. after the first spectroscopic observations on ut 2012 mar 27, the emission lines in the optical spectra became fainter relative to the continuum (on ut 2012 mar 28, apr 4 and 6), while the optical brightness was increasing slowly (the optical brightness changed gradually from 12 to 11 magnitudes in the v-band, see the american association of variable star observers (aavso) database, http:// http://www.aavso.org/lcg/). no emission lines could be observed (except hα emission with a p cygni profile), but many absorption lines of fe ii and neutral atoms such as na i (5890å) and o i (7773å) were detected clearly in the spectra obtained on apr 6 (t = 12 days). those absorption lines are indicative of lower ionization and the lower temperature conditions in the outflow of the nova. prominent c2 (swan) and cn (red system) absorption bands were detected on ut 2012 apr 8 (t = 14 days), as shown in figure 1. the obtained spectrum is similar to that of a carbon star. we also plotted the spectrum of tx psc (a well-known carbon star of spectral type n0;c6,2, with teff = 3030 k; lambert et al.,1986) for comparison. we also identified weak emission lines of hα, hβ, and fe ii. based on the substructure of the c2 swan band (∆v = -1) absorption, we could derive the isotopic ratio of carbon. figure 2 shows the spectrum of the nova, the modeled spectra of 12c12c, 12c13c, and 13c13c (with an excitation temperature of 4500k). the doppler shift of the nova spectrum has been corrected by using the relative velocity of the nova to the observer, estimated as 341 ± 87 km/s (this is derived from the velocities for hα and hβ emission peaks). the wavelengths of sub-peaks in this band cannot be explained by 12c12c only. clearly 213 m. nagashima et al. 12c13c and 13c13c contribute to form the absorption. 0.75 0.8 0.85 0.9 0.95 1 1.05 1.1 5400 5450 5500 5550 5600 5650 5700 5750 n o rm a liz e d f lu x wavelength [å] 13 13 c c 13 12 c c 12 12 c c apr. 08.7 figure 2: comparison between observed and modeled spectra of the c2 (∆v = -1) absorption band. the observed spectrum is shown by dashed line (the region influenced by the hg emission from the city-light, at ∼ 5460 å, has been removed). the modeled spectra for 13c13c, 12c13c and 12c12c are shown by the thin solid lines and the sum of those lines is shown by the thick solid line. for the modeled spectra, we assumed the rotational and vibrational temperatures of 4500 k and the isotopic ratio of 12c/13c = 4. the next observations were carried out on ut 2012 apr 16 (t = 22 days). in these observations, the c2 and cn absorption bands had already disappeared and strong balmer emission lines and fe ii lines were again prominent. at that time, the spectra were typical of fe ii-type classical novae. the other difference from the previous observations was the expansion velocity derived from hα, which had increased in comparison to that before molecular formation. this higher velocity is typical of the fe ii-type novae (williams 1992). thereafter, the spectra of this nova were not unusual for an fe ii-type nova, although we continued spectroscopic monitoring observations until ut 2012 may 26. the optical light curves showed a very slow decline after the visual maximum and a rapid drop (by about six magnitudes in the v-band) at around 90 days after the discovery (this nova can be classified as a ”slow” nova). the drop in the light curves may be caused by dust formation in the outflow of the nova. 3 results & discussion the equivalent widths of hα (and also of hβ) measured in our spectra were almost constant before and after the appearance of molecular absorption bands. however, optical light curves showed a small drop of ∼ 1 magnitude before and after molecular formation. figure 3 shows the optical and near infrared light curves taken from the database of the aavso and small and moderate aperture research telescope system (smarts), the color indices of (v − i) and the equivalent width of hα and hβ. for example, the c2 absorption could markedly affect band v-band magnitudes, while cn red absorption could also affect rand i-bands. this implies that extinction of the molecules in the outflow affected both continuum and emission lines from the nova. the molecular formation zone could be in the outer region of the outflow compared with regions emitting the continuum and/or emission lines in the nova. 1.2 1.4 1.6 1.8 2 v-i (aavso) v-i (smarts) -600 -500 -400 -300 -200 -100 0 0 10 20 30 40 50 60 70 80 90 100 110 120 hα hβ 5 10 15 20 25 v-mag. (aavso) v-mag. (smarts) i-mag. (aavso) i-mag. (smarts) k-mag. (smarts) figure 3: multi band light curve of v2676 oph (data from aavso and smarts), color indices (v −i) based on the aavso and smarts database, and equivalent widths of hα and hβ measured in our spectra. upper tick marks indicate days on which spectroscopic observations were performed. furthermore, molecular formation in v2676 oph is considered very rapid (within 2 days or less) and the existence of the molecular envelope was transient (it was present at most 9 days) at around the brightness maximum in optical. why was the appearance of both c2 and cn absorption bands so transient that they could be detected on apr 8 only? we considered that the outer region of the outflow became cool enough to form molecules, since the hard ultraviolet (uv) radiation from the pseudo-photosphere of the nova in the early phase was blanketed by an giron curtainh, and the iron ions could have absorbed uv radiation strongly (shore, 2008). this picture is consistent with the spectrum taken on apr 6, which was dominated by a continuum with absorption lines indicative of lower ionization conditions. the measurements of the color indices (v − i) obtained from the aavso and smarts database also showed a redder continuum for later periods after the discovery until the molecular bands appeared. this fact also supports the later lower color temperatures until molecular formation. however, as the envelope expanded and ejected materials rarefied (i.e., became more optically thin), hard uv radiation again increased in intensity. at this time, molecules would be destroyed 214 detection of diatomic molecules in the dust forming nova v2676 oph through photo-dissociation reactions caused by uv radiation. in support of this, the later spectra showed many emission lines from ionized species, such as fe ii. however, the molecular absorption bands may have disappeared due to some opacity effects. similar behavior in terms of cn formation was observed in dq her in 1934. the appearance of cn absorption bands immediately following the optical brightness maximum was transient, persisting for approximately 1 week only (sneden & lambert, 1975). the possible dust formation about 100 days after discovery was also similar to v2676 oph. theoretical studies of chemistry in the outflow of novae suggest that formation of even more complex molecules is possible (pontefract & rawlings, 2004; evans & rawlings, 2008). it has been demonstrated that a model atmosphere could reproduce both strong cn absorption bands in optical and the co emission band in the nearinfrared, as observed in some novae (hauschildt et al., 1994). although simple molecules might be destroyed by uv radiation, more complex molecules such as polycyclic aromatic hydrocarbons (pahs) (if they formed during the transient cool phase of the outer envelope), could survive and might act as nuclei for dust formation. indeed, pah emission was detected in this nova. we performed mid-infrared spectroscopic observation using a cooled mid-infrared camera and spectrometer (comics) mounted on the subaru telescope on ut 2013 june 20 (t = 452 days). the spectrum showed pah emission at 11.4 µm (and a hint of the emission line at 7.7 µm) on the smooth continuum that could be explained by amorphous carbon grains. 1 1.5 2 2.5 3 3.5 4 4.5 5 8 9 10 11 12 13 f lu x d e n si ty [ 1 0 -1 8 w /c m 2 /µ m ] wavelength [µm] pah figure 4: the mid-infrared spectrum of v2676 oph. references [1] mclaughlin, d.b. 1935, publ. aas, 8, 145. [2] geisel, s.l., kleinmann, d.e., low, f.j. 1970, astrophysical journal, 161, l101. doi:10.1086/180579 [3] wilson, o.c. & merrill, p.w., 1935, publications of the astronomical society of the pacific, 47, 53. [4] sanford, r.f. 1935, publications of the astronomical society of the pacific, 47, 209. [5] stoy, r.h. & wyse, a.b. 1935, publications of the astronomical society of the pacific, 47, 50. [6] antipova, l.i. 1969, soviet astronomy a.j., 13, 288. [7] sneden, c. & lambert, d. 1975, monthly notices of the royal astronomical society, 170, 533. doi:10.1093/mnras/170.3.533 [8] evans, a. & rawlings, j.m.c. 2008, in classical novae, 2nd edition. edited by m.f. bode and a. evans (cambridge astrophysics series, no. 43, cambridge: cambridge university press, 2008), p.308. [9] rudy, r. j., et al. 2012, electronic telegram no. 3103, central bureau for astronomical telegrams, international astronomical union (ed., green, d.). [10] nishimura, h. 2012, electronic telegram no. 3072, central bureau for astronomical telegrams, international astronomical union (ed., green, d.). [11] shinnaka, y., kawakita, h., kobayashi, h., naka, c., arai, a., arasaki, t, kitao, e., taguchi, g., ikeda, y. 2013, icarus, 222, 734. doi:10.1016/j.icarus.2012.08.001 [12] nagashima, m., et al. 2014, astrophysical journal letter, 780, 26 doi:10.1088/0004-637x/780/1/26 [13] arai, a. & isogai, m. 2012, electronic telegram no. 3072, central bureau for astronomical telegrams, international astronomical union (ed., green, d.). [14] lambert, d.l., gustafsson, b., eriksson, k., hinkle, k.h. 1986, astrophysical journal supplement series, 62, 373. doi:10.1086/191145 [15] williams, r.e. 1992, the astronomical journal, 104, 725. doi:10.1086/116268 [16] shore, s.n. 2008, in classical novae, 2nd edition. edited by m.f. bode and a. evans (cambridge astrophysics series, no. 43, cambridge: cambridge university press, 2008), p.194. [17] pontefract, m. & rawlings, j.m.c. 2004, monthly notices of the royal astronomical society, 347, 1294. doi:10.1111/j.1365-2966.2004.07330.x 215 http://dx.doi.org/10.1086/180579 http://dx.doi.org/10.1093/mnras/170.3.533 http://dx.doi.org/10.1016/j.icarus.2012.08.001 http://dx.doi.org/10.1088/0004-637x/780/1/26 http://dx.doi.org/10.1086/191145 http://dx.doi.org/10.1086/116268 http://dx.doi.org/10.1111/j.1365-2966.2004.07330.x m. nagashima et al. [18] hauschildt, p.h., starrfield, s., allard, f. 1994, in cool stars; stellar systems; and the sun; eighth cambridge workshop, astronomical society of the pacific conference series, vol. 64; proceedings of the 8th cambridge workshop (held in athens, georgia; october 11-14; 1993; san francisco: astronomical society of the pacific (asp); 1994; edited by jean-pierre caillault), p.705. 216 introduction observations results & discussion 148 acta polytechnica ctu proceedings 2(1): 148–151, 2015 148 doi: 10.14311/app.2015.02.0148 ss cygni revisited r. c. smith1, j. echevarŕıa2, j. v. hernandez2,3, p. szkody4 1astronomy centre, university of sussex, uk 2unam, mexico city, mexico 3physics & astronomy department, university of southampton, uk 4astronomy department, university of washington, usa corresponding author: r.c.smith@sussex.ac.uk abstract new spectroscopic and photometric observations of ss cygni, the brightest dwarf nova system, have been obtained, with the aim of mapping starspots on the surface of the secondary star. four nights of echelle spectroscopy in quiescence have been obtained using the 2.2-m telescope at san pedro martir (mexico) in august 2012 and another two nights at the 3.5-m telescope at apache point observatory, usa, in september 2012, but these data are still being reduced. simultaneous ccd photometry was also obtained at the two sites, and the mexican photometry was extended into the subsequent long outburst. this presentation reveals some interesting photometric behaviour in that outburst, but further data will be necessary before the nature of the behaviour can be determined. keywords: cataclysmic variables dwarf novae optical spectroscopy photometry individual: ss cyg. 1 introduction the canonical model of cataclysmic variable (cv) evolution requires secondaries to be magnetic, to allow magnetic braking to keep the star in contact with its roche lobe. magnetic activity should produce starspots, as shown by observations of rapidly rotating single stars. detection of starspots on cv secondaries provides evidence for magnetic fields, and roche tomography has been used to reveal spots on four systems: ae aqr (watson et al. 2006), bv cen, v426 oph (watson et al. 2007a, 2007b) and ru peg (dunford et al. 2012). the unusual nova-like system ae aqr is the brightest of these, and the others are at least two magnitudes fainter. ss cygni is the brightest dwarf nova, and in quiescence is only half a magnitude fainter than ae aqr. it is therefore an obvious candidate for study and highresolution echelle data were obtained in 2012 august (in mexico) and september (in the usa), with simultaneous photometry. the mexican echelle data were taken with the 2.2-m telescope at the san pedro martir (spm) observatory in baja california; the v photometry was taken with the spm 0.84-m telescope. the us echelle spectra were taken with the 3.5-m telescope at the apache point observatory (apo); simultaneous photometry was also obtained, using the nmsu 1-m telescope at apo (v ) and the uw 0.76-m at manastash ridge observatory (g,r,i). 2 spectra the us spectra were taken on two nights with the apo 3.5-m and have been fully reduced but not yet analysed. in figure 1 we show part of the average spectrum from the first night; the region shortward of 6750 å makes it clear that there are many absorption lines that can be used for roche tomography. four nights of data were taken with the spm 2.2-m; these spectra are still being reduced. figure 1: a portion of the average of 29 spectra of ss cygni taken with the echelle spectrograph on the 3.5-m apo on 20/21 september 2012. in addition to strong emission lines, there are many absorption lines suitable for roche tomography of the secondary. 148 http://dx.doi.org/10.14311/app.2015.02.0148 ss cygni revisited 3 photometry all the photometry has been fully reduced; here we just discuss the analysis of the mexican data. at the san pedro martir (spm) observatory in mexico, simultaneous v photometry with the 0.84-m telescope was taken to enable flux calibration of the echelle spectra being taken with the 2.2-m telescope. by good fortune, ss cygni went into outburst one day after the spectra had been taken, and our mexican colleagues kindly allowed us to continue photometric observing; the complete coverage is shown in figure 2. unfortunately, no useful data were obtained on the fifth night in outburst, which was clouded out; however, the mean level on that night does appear to be higher than on the previous nights, possibly lending support to the prediction of cannizzo (2013; see also cannizzo 2012) that long outbursts should have a precursor outburst at the beginning. subsequent aavso data confirmed that this was a long outburst. figure 2: v photometry of ss cygni in quiescence, rise and outburst, august 2012. we first analysed the quiescent data, looking for the dominant period. as expected, the only significant period to be found was the orbital period; a light curve folded on that period showed the characteristic double hump arising from ellipsoidal variation. there were four nights of useable outburst data. the data were first detrended, by removing linear trends from each individual night and then adding or subtracting suitable constants to bring all the nights to the same average magnitude. the starlink package period was used to search for the dominant period, using five different methods: string-length, minimum chisquared, lomb-scargle, fourier transform and clean. some of these methods produced no very useful results, and the clearest and most consistent results were obtained with the minimum chi-squared and lombscargle methods. the results below quote only the results from those two methods. figure 3: night 3 outburst data folded on a period of 0.35273 days and binned into 400 bins. note that the magnitude scale has brightness increasing downwards. analysing the entire data-set gave a best-fit period that was about 1.5% smaller than the orbital period, but the signal was not strong and this result is not thought to be significant. however, when the different nights were analysed separately a very different pattern emerged: each night had a different dominant period. the clearest result was for the 3rd night in outburst, where the two methods agreed on a period of 0.353 days. this is approaching the length of the data stream on that night (0.384 days), but is significantly shorter, and there is no sign of the data length in the period analysis. this period is also significantly longer than the orbital period of 0.27513 days. figure 3 shows the night 3 data binned and folded on a period of 0.35273 d. the best period on night 2 was close to the orbital period, with minimum chi-squared giving 0.2715 d and lomb-scargle giving 0.2769 d. however, the other two nights both gave significantly shorter best periods: 0.1433 d on night 1 and 0.1963 d on night 4. the lombscargle plots for all the nights are shown in figure 4 – note that the maximum power varies from night to night. thus, apart from night 2, the outburst data are not consistent with the orbital period, with two shorter periods and one longer one. there is a temptation to consider the night 3 data as evidence for a positive superhump – boneva et al. (2009) suggested that in outburst ss cyg has an elliptical disc. if so, ss cyg is not impossibly far off the standard psh,porb relation, and the orbital period of 6.6 h would be the longest on that relation. however, with a well-determined mass ratio q of 0.683 (bitner at al. 2007), it strongly violates the normal resonance criterion of q < 0.33. 149 r. c. smith et al. figure 4: lomb-scargle power plots for the four outburst nights. by far the strongest signal is on night 3, with a maximum power of ∼350. the maximum power on the other nights are: ∼50 (night 1), ∼200 (night 2) and ∼100 (night 4). the dominant periods on nights 1 to 4 are 0.1433 d, 0.27-0.28 d, 0.35273 d and 0.1963 d respectively. the horizontal axis is frequency, in cycles/day. furthermore, the amplitude is low (∼0.08 mag), the superhump excess (28%) is rather large, and two of the other nights suggest negative superhumps (although these have even larger differences from the orbital period, at 48% and 29% smaller, and have similarly small amplitudes). is there a better explanation? one possibility was raised by bisikalo (2013; see also zhilkin & bisikalo 2010), who showed models where fluctuations in the accretion rate onto the white dwarf, caused by variations in the generation of magnetic field in the disc, produced brightness variations on various timescales. perhaps we are seeing evidence in our data for similar brightness variations, which might be stochastic in nature. 4 conclusions the spectral data we have obtained for ss cygni appear to be good enough for us to be able to map starspots on the surface of the secondary component. the photometric data in quiescence appear to show ellipsoidal variations on the orbital period. however, the photometric data in early outburst are generally not consistent with the orbital period. the data from the 1st and 4th nights have shorter periods (negative superhumps?) while the data from the 3rd night show a longer period (a positive superhump in the longest-period system so far?). alternatively, we may have found evidence for variations in the accretion rate onto the white dwarf, caused by magnetic effects in the disc. whatever the explanation, it would be worth monitoring ss cygni intensively during its long outbursts, to see whether this peculiar behaviour is repeated. it turns out that the kind of photometric data we obtained during the long outburst in august 2012 is rare: not many people have carried out time-resolved observations of the long outbursts in dwarf novae and so the nature of the behaviour during these outbursts is still 150 ss cygni revisited quite uncertain (cf. cannizzo 2012, 2013). in order to test whether any changes are periodic or stochastic, it will be necessary in future to monitor as many of these long outbursts as possible, in ss cygni and in similar dwarf novae. it is hoped to do this by organising an international campaign involving both the amateur community and the many robotic telescopes scattered around the world. by the time these data are available, it is hoped that predictions of the magnetic effects in the disc, by bisikalo and others (e.g. bisikalo 2013), will have reached the stage where they may be compared in detail with the data to discover whether the model is compatible with observations. acknowledgement we thank leslie hebb for her painstaking reduction of the apo spectra, our mexican colleagues for allowing us to observe ss cyg in outburst, phil charles for comments on the outburst photometry and tom marsh for help with software. we acknowledge the use of the period program from the starlink package, supported at the joint astronomy centre in hawaii, and of the aavso light curve generator. references [1] bisikalo, d.: 2013, these proceedings. [2] bitner, m.a., robinson, e.l., behr, b.b.: 2007, apj 662, 564. doi:10.1086/517496 [3] boneva, d. et al.: 2009, astronomy reports 53, 1004. doi:10.1134/s1063772909110055 [4] cannizzo, j.k.: 2012, apj 757, 174. doi:10.1088/0004-637x/757/2/174 [5] cannizzo, j.k.: 2013, these proceedings. [6] dunford, a., watson, c.a., smith, r.c.: 2012, mnras 422, 3444. [7] watson, c.a., dhillon, v.s., shahbaz, t.: 2006, mnras 368,637. [8] watson, c.a., et al.: 2007a, an 328, 813. [9] watson, c.a., et al.: 2007b, mnras 382, 1105. doi:10.1111/j.1365-2966.2007.12173.x [10] zhilkin, a.g., bisikalo, d.v.: 2010, astronomy reports 54, 840. doi:10.1134/s1063772910090088 discussion david buckley: i am puzzled by the superhump excess � for tx col. is there a reference for the superhump period? robert smith: i took the superhump period and excess for tx col from montgomery (2009, apj, 705, 603), who quoted retter et al. (2005, assl, 332, 251). the retter et al. conference paper is suggestive but not definitive. christian knigge: since you typically see only about one cycle of your periods in each epoch, is it not possible that the variability is stochastic rather than periodic? robert smith: in some cases we have two cycles. but more generally, yes, to call them periodic changes is a bit speculative. however, there is a clear single peak in the lomb-scargle power spectrum on each night, and they are certainly at different frequencies on different nights. the night-to-night changes may be stochastic, as you suggest, but there is definitely something interesting happening. 151 http://dx.doi.org/10.1086/517496 http://dx.doi.org/10.1134/s1063772909110055 http://dx.doi.org/10.1088/0004-637x/757/2/174 http://dx.doi.org/10.1111/j.1365-2966.2007.12173.x http://dx.doi.org/10.1134/s1063772910090088 introduction spectra photometry conclusions acta polytechnica ctu proceedings doi:10.14311/app.2017.8.0017 acta polytechnica ctu proceedings 8:17–19, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app microscopic evaluation of the quality of dental replacement františka pešlováa, b, ∗, daniela koštialikováa, richard veselýc, andrej dubeca, maxim puchninb a faculty of industrial technology in púchov, alexander dubček university of trenčín, i. krasku 491/30, púchov, slovakia b czech technical university in prague, faculty of mechanical engineering, karlovo nám. 13, 121 32 prague 2, czech republic c czech technical university in prague, faculty of mechanical engineering, department of enterprise management and economics, karlovo nám. 13, 121 32 prague 2, czech republic ∗ corresponding author: abstract. the permanent tooth replacements including metal-ceramic crowns are a convenient solution for the renewing of the original function of the whole set of teeth as well as of the natural appearance. development and preparation of suitable tooth replacement presents a real challenge for dental engineers as the replacement has to meet all the conditions and requirements of the dental medicine as well as patient's needs and wishes. the preparation of the metal-ceramic crown is a sophisticated process and the dental engineer has to prepare always a unique replacement on demand. in dental medicine there is a wide spectrum of inorganic and organic materials used for manufacturing of dental replacements. each of the material has specific properties leading to distinct applications. besides the material properties, the attention has to be paid to the aesthetic function and biocompatibility of the material to ensure the complete restoration of the whole set of teeth. keywords: metal-ceramic crown, dental replacement, microstructure. 1. introduction the presented paper focuses on investigation of the randomly selected tooth replacements which were no longer suitable to be used in mouth cavity in a satisfying and predetermined way. from the aspect of materials engineering, the mentioned dental condition or phenomenon is based on the critical state of the dental replacement and it can be the cause of the occurrence of many other critical states. the given critical states lead to degradation of tooth replacements and it can even end in tooth replacement's fracture causing the functional and aesthetic defect as well. [1–4]. the main scope of the performed microscopic observations was to identify the initiation stimulus for rupture of a tooth replacement – the child patient's crown, which was used as tooth replacement after the injury. in relation to the given metal-ceramic crown (of the type seen in the 1), there was the rupture of the ceramic layer and its subsequent delaminating. the crown was supplied without any specific information (producer, chemical composition, life duration, etc.) 2. materials the metal-ceramic crown is a complex composite material with chemical composition given in the table 1 and 2. the identification of the phase composition was evaluated using scanning electron microscope (sem) jsm-7600f (jeol, jp) with energy dispersive spectroscopy (eds) x-max 50 mm2 (oxford instruments, gb) was used.the metallographic preparation of samples was performed in a standard and normalized way in order to carry out the microscopic investigation of the microstructure of used materials in their crosssections (fig. 2 ). metallographic preparation was made in standard way – selection of the material from the defect and from the intact location followed by grinding, polishing and etching. according to the fact mentioned herein before, the precise investigation of ceramics and metal microstructure were carried out. the detailed images (fig. 3) revealed the sites with the altered roughness. the occurrence of cavities was identified in sites where the metal was in contact with the ceramic layer and it led to the rupture of the ceramic layer. based on the microscopic observations, the penetration of ceramic material into the irregular areas of metal originating from the manufacturing process has been identified. that led to the deteriorative effect in relation to the metal – ceramics contact (fig. 4). the microscopic observation of the metal alloy quality revealed that the low-quality heterogeneous metal alloy was used for manufacturing of selected metalceramic crown which caused defects in contact layer between metal and ceramic material during manufacturing. besides the distribution of individual phases, 17 http://dx.doi.org/10.14311/app.2017.8.0017 http://ojs.cvut.cz/ojs/index.php/app f. pešlová, d. koštialiková, r. veselý et al. acta polytechnica ctu proceedings weight % o si sn k co al cr zr ti na sample nr.1 – ceramic 34,2 14,8 8,5 6,4 6,3 5,2 2,9 2,7 2,6 2,2 table 1. chemical composition of the ceramic part of the crown obtained by edx analysis weight % co cr w o si v s al k sample nr.1 – metal 57,2 22,7 6,9 2,1 1,4 0,6 0,5 0,5 0,3 table 2. chemical composition of the metal part of the crown obtained by edx analysis figure 1. metal-ceramic crown, detail. figure 2. metal-ceramic crown detailed image. figure 3. rupture in the site of the metal-ceramic material contact. figure 4. penetration of ceramic material. the low quality of materials was confirmed by chemical composition. the character of the microstructure (fig. 5) exhibits the imperfect primary crystallisation of co-cr alloy with the occurrence of undesired phases. [5–8]. 3. conclusion based on the microscopic evaluation, it can be concluded that individual microstructure of the metal alloy and contact layer between metal alloy and ceramic material indicates that evaluated metal alloy was not suitable for manufacturing of tooth replacement. the most significant evidence was obtained from the contact layer between metal and ceramic material where cavities were observed. the cavities are a source of oxidation which can lead to further degradation of the metal alloy. with the reference to the obtained results, we can come to the conclusion that the low quality materials used for manufacturing process of the metal-ceramic crown had caused defects that led to the degradation of the replacement. moreover, there was a danger of damaging other teeth as well as danger of ingestion or inhaling parts of metal or ceramic material. the work is showing an example that may occur in practice thus a regular inspections of patients tooth replacements are recommended. 4. results regulatory requirements for control samples preparation during the manufacturing of the replacements should be initiated despite the fact that preparation of samples for potential defect detection control may not be economical manufacturing process even though it is in the best interest of the patients. 18 vol. 8/2017 microscopic evaluation of the quality of dental replacement figure 5. microstructure of co-cr alloy . acknowledgements this work was supported by the slovak grant agency kega 006tnuad-4/2014, ministry of education youth and sport of the czech republic program npu1, project no lo127 references [1] y. wang. bioadaptability: an innovative concept for biomaterials. journal of materials science & technology 32(9):801 – 809, 2016. bioadaptation of biomaterials, doi:http://dx.doi.org/10.1016/j.jmst.2016.08.002. [2] h. hubálková, j. krňoulová. materiály a technologie v protetickém zubním lékařství. galén, praha, 2009. [3] r. g. craig, j. m. powers. restorative dental materials. mosby, st. louis, missouri, 11th edn., 2002. [4] h. hubálková, j. charvát, t. dostálová. faktory ovplyvňujúce životnosť fixnej zubnej náhrady. progresdent, 5th edn., 2004. [5] j. giacchi, c. morando, o. fornaro, h. palacio. microstructural characterization of as-cast biocompatible co-cr-mo alloys. materials characterization 62(1):53 – 61, 2011. doi:http://dx.doi.org/10.1016/j.matchar.2010.10.011. [6] j. giacchi, o. fornaro, h. palacio. microstructural evolution during solution treatment of co-cr-mo-c biocompatible alloys. materials characterization 68:49 – 57, 2012. doi:http://dx.doi.org/10.1016/j.matchar.2012.03.006. [7] j. e. lemons, f. misch-dietsch, m. s. mccracken. biomaterial for dental implants. in dental implant prosthetics, chap. 4, pp. 66–94. mosby, st. louis, missouri, 2015. [8] j. b. suzuki, d. lynn, l. d. terracciano-mortilla, c. e. misch. maintenance of dental implants. in dental implant prosthetics, chap. 34, pp. 964–981. mosby, st. louis, missouri, 2015. 19 http://dx.doi.org/http://dx.doi.org/10.1016/j.jmst.2016.08.002 http://dx.doi.org/http://dx.doi.org/10.1016/j.matchar.2010.10.011 http://dx.doi.org/http://dx.doi.org/10.1016/j.matchar.2012.03.006 acta polytechnica ctu proceedings 8:17–19, 2017 1 introduction 2 materials 3 conclusion 4 results acknowledgements references 103 acta polytechnica ctu proceedings 2(1): 103–106, 2015 103 doi: 10.14311/app.2015.02.0103 why a new code for novae evolution and mass transfer in binaries? g. shaviv1, i. idan1, n. j. shaviv2 1department of physics, israel institute of technology, haifa, israel 2racah institute of physics, the hebrew university, jerusalem, israel corresponding author: gioras@physics.technion.ac.il abstract one of the most interesting problems in cataclysmic variables is the long time scale evolution. this problem appears in long time evolution, which is also very important in the search for the progenitor of sn ia. the classical approach to overcome this problem in the simulation of novae evolution is to assume: (1) a constant in time, rate of mass transfer. (2) the mass transfer rate that does not vary throughout the life time of the nova, even when many eruptions are considered. here we show that these assumptions are valid only for a single thermonuclear flash and such a calculation cannot be the basis for extrapolation of the behavior over many flashes. in particular, such calculation cannot be used to predict under what conditions an accreting wd may reach the chandrasekhar mass and collapse. we report on a new code to attack this problem. the basic idea is to create two parallel processes, one calculating the mass losing star and the other the accreting white dwarf. the two processes communicate continuously with each other and follow the time depended mass loss. keywords: nova modeling thermonuclear runaways. 1 introduction the classical prediction or identification of a sn ia progenitor as a wd in a compact binary system is usually based on the calculation of a single thermonuclear flash or at most few and how much mass the wd gains or loses in such a flash. extrapolations of the behavior of binary systems based on single or few thermonuclear flashes are not expected to be reliable. prialnik and kovetz (1995) were the first to simulate numerically a rather long series of thermonuclear runaway, up to a 1000. these authors solved in this way the question of the initial conditions assuming that after so many flashes, the wd converges to periodic behavior. idan et all (2013) carried a similar calculation for a high accretion rate and the results were not similar, nor were they strictly periodic. in both calculations the rate of accretion was constant in time and many flashes were calculated. however, the behavior of the mass losing star and its response to mass loss are not uniform in time so that the assumption of constant mass loss (and accretion at a constant rate on the wd) is not justified. for a star of radius r undergoing mass loss dr dt = ( ∂r ∂t ) ev + ∂r ∂m ṁ. (1) the index ev means change of r due to normal secular stelar evolution. one usually assumes that ( ∂r ∂t ) ev � ∂r ∂m ṁ. (2) but then r − rroche 6= constant and clearly, the expression for mass loss due to roche lobe overflow yields a non constant rate of mass loss, as this expression depends on r(t). note that if we assume that dr/dt = 0, namely an equilibrium or a steady state and we neglect the time variability of rroche, then ṁ = − ( ∂r ∂t ) ev / ∂r ∂m � 1, (3) in units of solar radius per solar mass. hence this expression does not yield observable values and does not imply that ṁ(t) = const. hence, this assumption is unacceptable. webbink (1977) evaluated the mass derivative of the roche radius and obtained: ( ∂ ln rroche ∂ ln m ) m̄,j = f(q) (4) where q = ln(m + 0.005m∗)−0.5 ln p−ln(p + 1016) has to do with the mass shell division of the calculation. webbink also assumed that a = ( d ln r d ln m ) t ≈ d ln r d ln m � d ln rroche d ln m (5) 103 http://dx.doi.org/10.14311/app.2015.02.0103 g. shaviv, i. idan, n. j. shaviv where a is the adiabatic constant and the inequality holds for all ṁ. 2 mass loss rate there are several empirical expressions for the mass loss. webbink (1977) for example, assumed that ṁ = −λ ( r−rroche r )2 with λ = const (6) provided r − rroche > 0, while ritter (1988) wrote that ṁ = ṁ0 exp (r−rroche) hp for r−rroche > 0. (7) here m0 is a constant to be evaluated from the geometry of the roche lobe while hp is the pressure scale height near the l3 point. it is clear that if these expressions are valid, then a change caused by a change in the radius of the star affects the accretion rate (and the nova long time evolution). 3 time scale involved the reaction of the donor star depends on three time scales. (a) the kelvin-helmholtz-ritter time scale given by: τkhr = gm2donor rldonor ( ∆m mdonor ) , (8) where ∆m is the mass affected by the mass loss perturbation. (b) the accretion time scale τacc = ∆mflash ṁ (9) where ∆mflash is the accreted mass at which the thermonuclear runaway takes place (of the order of 10−5m�) and (c) the dynamic time scale is given by: τdyn = √ 3 4πgρ̄ . (10) here ρ̄ is the mean density of the star. only mflash, the mass at which the nuclear flash occurs, depends (slightly) on the mass of the wd. the dynamic time depends on the entire star and τkhr depends on the outermost mass-shell involved. the interplay between these three time scales controls the phenomenon and it varies with the rate of mass loss. consequently, we calculated the hydrodynamic and thermodynamic evolution of the star under the condition that the accretion rate is given (by the parameters of the binary system). as the present calculation is carried out irrespective of the mass of the accretor, we cannot evaluate the accretion rate but have to impose it as given. 4 the dynamic behavior of the donor our goal is to investigate the dynamic response of the donor star to mass loss. we assume spherical symmetry and solve the full hydrodynamic equations of the donor and evaluated the requested derivatives. in figure 1 we see how dr/dm behaves in the case of low accretion rate (10−10m�/yr). the donor is a main sequence star of 1.25m�. 0 2000 4000 6000 time(years) 2x105 4x105 6x105 8x105 106 0 – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ figure 1: the derivative dr dm in units of m�/τkhr for an imposed mass loss of 10−10m�/yr. in this case τflash ≈ 105yr. we realize that neither dr/dt nor (∂r/∂m)ṁ are constant and one should not expect the accretion rate to be constant (in time) either. even during the period of mass building for a single flash, the accretion rate is not constant. – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ 0 0.5 1 1.5 2 2.5 3 1013 1012 1011 1010 109 108 time(years) figure 2: the derivative dr/dm in units of m�/τkhr for an imposed mass loss of 10 −6m�/yr. in this case τflash ≈ 10yr. the results for all accretion rates are collected and summarized in figure 3. we see that in all cases, irrespective of the accretion rate, the time dependence of the derivative is given by: dr dm = 14.07 t2.35 + �(ṁ) for all ṁ, (11) where �(ṁ) is a constant in time which depends on ṁ. 104 why a new code for novae evolution and mass transfer in binaries? 10-6 10-5 10-4 10-3 10-2 10-1 100 101 102 103 1014 1013 1012 1011 1010 109 108 107 106 105 104 103 102 – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ time (years) m*=1.25msun figure 3: the derivative dr/dm in units of m�/τkhr for all accretion rates calculated here (for a ms star of mass 1m�). moreover, time dependence of the derivatives tend for sufficiently long times, to an asymptote. we find that: for t � �(ṁ) dr dm → 14.07 t2.35 , (12) namely, all results (for a given mass of the donor) converge for long times to an asymptote. we do not know at the moment how this asymptote changes with the mass of the donor. at the same time we can write for the roche lobe radius (eggelton, 1984) that: rroche ≈ 0.49q2/3 0.6q2/3 + ln(1 + q1/3) where q = mwd + ṁt md − ṁt (13) assuming conservative mass loss. hence for sufficiently small t (at the beginning) we have that: ∣∣∣∣ drroche dm ∣∣∣∣ � ∣∣∣∣ drdonor dm ∣∣∣∣ (14) the thermodynamic state is shown in figure 4 time(years) – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�9m�/yr 10�10m�/yr @r @m ⇣ r� m� ⌘ – 3 – 10�11m�/yr (lph�lnuc) lnuc @r @m ⇣ r� m� ⌘ – 2 – where index ev means change of radius due to stellar evolution and @r/@m means change of radius due to change in the total mass due to mass loss from the surface. this equation does not expose the full processes taking place in the star and assumes that @r/@m) does not depends on ṁ. moreover. the derivative @r/@m) is usually calculated assuming the star is in thermal equilibrium and in strict hydrostatic equilibrium, namely the effect of ṁ on the process is ignored. if we do neglect these effect then we may ask what happens when dr/dt were to vanish. in this case we could write that ṁ = � � @r @t � ev @r @m . (2) this however, yields extremely small ṁ and implies that the assumption of ’steady state’, namely dr/dt = 0, is untenable. the thermal relaxation of the donor star in response to the mass loss of mass �m is given by ⌧khr = gm 2d rld ✓ �m md ◆ (3) where ld is the luminosity of the donor and md its mass. the accretion time scale is given by ⌧acc = mf lash ṁ (4) where mf lash is the accreted mass which leads to a thermonuclear runaway and ṁ is the accretion rate generated by the donor envelope. we assume that of the mass lost by the donor a fraction 1 � ⌘ is lost and does not reach the accreting star. we first ask when the thermal time scale of the donor is of the same order and the accretion time, namely gm 2 rld �m md = ⌘�mmw d ṁ (5) or: ṁ mw d = rld gm 2d (6) and hence the critical accretion rate is: (7) 2. references 10�6m�/yr 10�7m�/yr 10�8m�/yr 10�9m�/yr 10�10m�/yr 102 101 100 10-1 10-2 10-3 10-4 10-7 10-6 10-5 10-4 10-3 10-2 10-1 100 101 102 103 – 3 – 10�11m�/yr (lph�lnuc) lnuc @r @m ⇣ r� m� ⌘ figure 4: the time to reach thermodynamic equilibrium. we see that stars with mass loss rate smaller than 10−9m� tend to thermal equilibrium. the timescales to reach the thermal equilibrium vary. accretion rates higher than 10−9m� diverge, namely they become unstable and runaway. the rate of accretion drives the star out of thermal equilibrium to be never restored. this fact should be taken into account in evaluating the mass loss rate from the donor. 5 conclusions first conclusion: it is not justified to assume that τkhr is negligible. second, polytropic estimates are nice and simple, but wrong (motl et al.2002) the mass loss is not constant in time. the mass loss does not start suddenly and reaches the assumed value gradually. there is no fixed period between eruptions. the system can have n eruptions with an almost constant time interval and then pause and let the donor recover on a kelvin-helmholtz-ritter time scale of its envelope. during this time the wd may relax to a new state. the accretion rate changes in time and affects the evolution of the nova. a nova calculation must include the evolution of the donor and the accretion rate. an important element in the evolution of nova is the time variability of the accretion rate. 0 2x1013 4x1013 6x1013 8x1013 1x1014 -0.0014 -0.0012 -0.001 -0.0008 -0.0006 -0.0004 -0.0002 0 time (seconds) m as s lo st fr om d on or (s ol ar m as se s) figure 5: the time depended mass loss from the donor. 6 discussion and conclusions a significant part of present day interest in nova includes the cases that may become progenitors of sn ia. in this case the behavior of the binary system is followed through a single or few thermonuclear runaways and then the result is extrapolated over 6-7 orders of magnitudes. the mass accreted or lost is of the order of 10−7m�. the initial mass of the wd is of the order of 1m� and if a conservative mass transfer is assumed this means a huge extrapolation in the behavior of the accretor and the the donor. we conclude that the evolution of the two stars must be followed simultaneously. we developed a code which 105 g. shaviv, i. idan, n. j. shaviv does just that. two processors are created, each devoted to a star. thus the donor star is calculated on one cpu and the accretor on a second cpu. the two processes, which may run on different computers or on a computer with more than one cpu, communicate with one another via an open gate. the communication can take place at fixed time intervals or whenever the conditions on one star deviate significantly and an update is due. 0 2x1013 4x1013 6x1013 8x1013 1x1014 0 5x107 1x108 1.5x108 2x108 2.5x108 time (seconds) l nu c / l su n l nu c/ l su n figure 6: the flashes of nuclear energy as a function of time. the code is in the debugging phase but the results seem to justify the claims present here. we find that the accretion rate is not constant in time and it stops when the radius of the donor shrinks below the roche lobe radius. in figure 5 we show one such example. the periods of no accretion appear as horizontal section (no change in mass lost). the period of mass loss appear as decreasing lines. note that these parts of the curve are straight lines at the beginning but not later. hence the rate of mass loss changes even during accretion. the recovery time depends on the mass loss rate and the τkhr of the donor. the resulting nuclear flushes are shown in figure 6. we see that the picture of flashes at a constant rate is correct only of a couple of flashes and the flashes come in groups. the time between the groups, the relaxation time of the donor increases gradually. acknowledgement this research has been supported by the israel science foundation, grant 1589/10. shaviv giora is delighted to thank dr. ilan shaviv in setting the code. references [1] eggleton,p.p.1983,apj.,268,368. doi:10.1086/160960 [2] idan, i., shaviv, n.j. & shaviv, g. , 2013, mnras, 433, 2884 doi:10.1093/mnras/stt908 [3] kolb, u. & ritter, h. 1990. a&a, 236, 385. [4] motl, p.m., tohline, j.e. & frank, j., 2002, apjs, 138,121. doi:10.1086/324159 [5] prialnik, d. & kovetz, a., 1995, apj, 445, 789. [6] ritter, h. 1988, a&a, 202,93. [7] webbink, r. f. , 1977, apj, 211, 881. 106 http://dx.doi.org/10.1086/160960 http://dx.doi.org/10.1093/mnras/stt908 http://dx.doi.org/10.1086/324159 introduction mass loss rate time scale involved the dynamic behavior of the donor conclusions discussion and conclusions acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0005 acta polytechnica ctu proceedings 12:5–9, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app model car transport system – modern its education tool karel bouchner∗, alina mashko department of vehicles, ctu in prague, horská 3, praha 2, 120 00, czech republic ∗ corresponding author: bouchkar@fd.cvut.cz abstract. the model car transport system is a laboratory intended for a practical development in the area of the motor traffic. it is also an important education tool for students’ hands-on training, enabling students to test the results of their own studies. the main part of the model car transportation network is a model in a ratio 1:87 (ho), based on component units of faller car system, e.g. cars, traffic lights, carriage way, parking spaces, stop sections, branch-off junctions, sensors and control sections. the model enables to simulate real traffic situations. it includes a motor traffic in a city, in a small village, on a carriageway between a city and a village including a railway crossing. the traffic infrastructure includes different kinds of intersections, such as t-junctions, a classic four-way crossroad and four-way traffic circle, with and without traffic lights control. another important part of the model is a segment of a highway which includes an elevated crossing with highway approaches and exits. keywords: model car transportation system (model), motor traffic, intelligent transport system (its), smart city. 1. introduction with a constant growth of people mobility, expansion of cities infrastructure, increase in traffic flow, traffic density in the cities, the efficient traffic management systems are rather crucial. telematic systems are being used for traffic monitoring and control. the road traffic is affected by a number of parameters such as weather conditions, infrastructure, human behavior, technical state of a vehicle etc. telematic systems are used in city planning, transport management, parking solutions, operation of logistics and municipal services, organization of urban space, implementation of sustainable energy sources. it can as well as be comprised under an umbrella term, the smart city initiative, and it is approached from the view of different disciplines including technical, economic, humanitarian, legal etc. [1]. traffic modeling is important in the process of traffic systems implementation. the preliminary system simulation and testing in laboratory conditions with implementation of real traffic data is safe and provides feasible results that can be used for timely corrections and improvement in traffic management schemes. there are many advantages of performing traffic studies in safe laboratory conditions. these include possibility of simulating the enforcement of services work in emergency situations, simulating dangerous human behavior in traffic with a high density, dangerous road sections or on high-speed roads, planning safer routes for dangerous goods transportation, efficient city traffic planning including city municipal transport, parking areas, better overview of city safety in general etc. rather crucial for smart city technologies is testing newly developed telematic systems on real traffic data which can be realized within the suggested model[2, 3]. the smart city solutions are to take into account the existing city architecture while dealing with efficient energy flows[4]. the mathematical (or general) models of such complex systems provided with the use of tools of virtual reality to the users are nowadays very popular and widely applied. however, the real interpretation of the models is still for many people straightforward for understanding and working with. the real models are not only convenient for demonstration and illustration purposes, but they are also especially convenient for education and training. the model car transport system described herein presents a laboratory with a physical model of infrastructure and is intended for research and analysis of road traffic monitoring and management and is realized at department of vehicle technology at faculty of transportation sciences at ctu in prague. a similar project has been realized for rail traffic model within the laboratory of transport technology[5] and is running at department of transport telematics of ctu, prague. 2. system overview the basis of the model car transportation network is a model in a ratio 1:87 (ho), based on component units of faller car system[6], e.g. vehicles, traffic lights, carriageway, parking spaces, stop sections, branch-off junctions, sensors and control sections (see figure 1[6]). the faller car system was initially designed to complement railway models. in our case it is used the other way around. the fundamental car model is 5 http://dx.doi.org/10.14311/app.2017.12.0005 http://ojs.cvut.cz/ojs/index.php/app karel bouchner, alina mashko acta polytechnica ctu proceedings figure 1. some components of the model. source: [6]. figure 2. city (left) and village (right) infrastructure. supplemented with a small railway that among other things includes two railway stations and a railway crossing. it enables simulation of real traffic situations on intraand extravilan communications with a level crossing with a rail road and a section of highway (see figures 2 through 5 for reference). see the general schematic representation of the whole model on figure 6. 2.1. technical description of the model road in the model is realized with the help of wooden plates. the technology that controls vehicles’ lateral movement is allocated within the road surface (wire rail) and under the surface. model vehicles are equipped with an electric motor, a steering apparatus, a rechargeable battery, an on/off switch on the bottom of the car to start it and a reed sensor that responds to magnet effect. permanent magnet tip is attached in a flexible way under the road so that an approaching vehicle with steering mechanism in front responds to wire magnetism which is keeping vehicles on track. 2.1.1. physical layer of the model infrastructure physical part of the model is represented by the road that is built up with the help of wooden plate elements figure 3. intravilan connection. figure 4. highway section with a trumpet-type interchange. fixed on foam plastic pieces of various constructions for modeling of different terrain levels. a wire rail is embedded into the wooden road along its length. control sections are placed under the road (see figures 7 through 9 for physical layer elements for reference). sensors are used for activation of functional elements, for example stop sections, parking spaces and branch-off junctions. they contain reed contact. sensors embedded in the road are activated by magnets on the vehicles when the vehicle goes over the sensor. the sensor gives precise feedback on the traffic control. this signal activates the control of a functional element. 2.2. traffic control • traffic node control – intersection the stop section (see figure 10) is needed to initiate vehicle full stop at road junctions (intersections and level intersections), bus stops, parking slots etc. it consists of an electromagnet embedded in the roadway. with power, its magnetic field will switch off the vehicle’s motor power via the reed sensor. for example, stop sections together with traffic lights and hw control section can control traffic on a classic four-way crossroad in a completely automated way (see figure 11). • traffic node control – navigation and routing the branch-off junction (see figure 12) is used for turning of vehicles to the right or left. when activated, the magnetic field of the branch-off junction diverts a vehicle onto a second, branching contact wire. this is done through the magnet on the vehicle’s steering slider. 6 vol. 12/2017 model car transport system figure 5. supermarket with parking. figure 6. general schematic representation of traffic real model. • parking control permanent magnet is also applied in case of longer stops such as parking (see figure 13). when the parking space is activated, an integrated electric coil briefly interferes with this magnetic field. this closes the reed sensor in the vehicle and supplies power to the motor. the vehicle starts moving. the traffic infrastructure includes different types of intersections including t-junctions, a classic fourway crossroad with traffic lights and a four-way traffic circle, bus stops etc. an important part of the model is a segment of a highway which includes an elevated crossing with highway entries and exits. the highway has two lanes in each direction. such an arrangement corresponds to common real construction of highways (according to standard [7], except for a few parts where accommodation was needed due to space constrains). the two-lane arrangement enables to simulate real traffic situations, e.g. when one vehicle enters the highway and the second one is approaching to the highway at the same time, or during overtaking of vehicles on the figure 7. road with metal wire, with control element. figure 8. physical part of road construction in different sections (left – four-way highway, right – elevated intersection) and model vehicles. highway. a disadvantage of the technology is that a movement of model vehicles does not fully correspond to a behavior of real vehicles. model vehicles move by a constant speed. the speed depends only on a status of a rechargeable battery. when the battery is fully charged the speed is the highest. during its discharging, the speed of model vehicle decreases. so far, model vehicles are neither able to start and stop smoothly, nor to vary their speed. in addition, now, their speed does not correspond to a type of real vehicles. 2.3. model operation and control there are three different ways to control the whole model: • manual control an operator activates the control of functional elements manually, using push buttons and switches. model vehicles move according to the operator’s immediate actions. • preset automatic control the functional elements are controlled by associated control sections. the operation of the control sections can be either fully automatic, depending only on their initial setting, or depending on their initial 7 karel bouchner, alina mashko acta polytechnica ctu proceedings figure 9. examples of three types of control sections (from left to right): parking space, stop section and branch-off junction – an alternative to faller elements. figure 10. stop section. source: [6]. setting in combination with signals from the associated sensors. model vehicles move according to their immediate status of functional elements controlled by the corresponding control section. operation of individual control sections is not synchronized. • program control the functional elements are controlled by an associated pc-standard module and expansion modules. the pc-standard module is connected to a pc with an appropriate sw application and interconnected with expansion modules. also sensors are connected to the modules. an operator creates on the pc programs of different traffic situations for the whole model, using the sw application. the sw application also utilizes signals from sensors for its activity. modules operate according to the running program. model vehicles move according to their immediate status of functional elements controlled via corresponding modules by the running program. figure 11. example of intersection control. source: [6]. figure 12. branch-off junction. source: [6]. 3. further perspective the current model shall be used as the basis platform for future smart systems implementation, programming based on real traffic data, or for simulation of real traffic and driver behavior. it brings several challenges for the development of software applications for intelligent transport systems, implementations of more traffic data collectors such as a camera or road section detectors and other traffic surveillance and control systems[8]. the system may serve for students’ training in a wide range of disciplines applied in transport studies and research, namely, traffic management and control, city planning and infrastructure design, software development, driver behavior and humanmachine interaction just to name a few. the current project has a goal to design a fully automatically operated system with simulation of vehicle-to-vehicle (v2v), infrastructure-to-vehicle (i2v or v2i) commu8 vol. 12/2017 model car transport system figure 13. parking space. source: [6]. nication in a way so that fully autonomous vehicles could drive in real conditions. the video-based data from the cameras that will be installed in the system will provide the traffic information for decisions to be taken by robot drivers of individual vehicles. the cameras are to be installed in the control sections and on the vehicles for strategic route planning as well as immediate reaction to the traffic in front, correspondingly. the front cameras can also be used for manual vehicle control with remote driver. thus, it is planned to simulate possible future scenarios of vehicle operation in smart cities[9]. 4. conclusion the model car transport system is one of important tools, assisting students of faculty of transportation science in their study. it facilitates understanding of matters of the intelligent transport system and the smart city. finally, it enables a comprehensive application of student theoretical knowledge in a wide range of transportation theory topics, in informatics, electronics, control theory. it is an appropriate tool for software and model hardware testing of various its applications and tools applicable in smart city projects. references [1] m. lom, o. přibyl, m. svítek. industry 4.0 as a part of smart cities. conference paper. smart cities symposium prague 2016. doi: 10.1109/scsp.2016.7501015. [2] l. galán-garcía, g. aguilera-venegas, p. rodrígues-cielos. an accelerated-time simulation for traffic flow in a smart city. journal of computational and applied mathematics. volume 270, november 2014, pages 557-563. [3] z. li, m. shahidehpour. deployment of cybersecurity for managing traffic efficiency and safety in smart cities. the electricity journal. volume 30, issue 4, may 2017, pages 52-61. [4] c. navarro, m. rca-riu, s. furió, m. estrada. designing new models for energy efficiency in urban freight transport for smart cities and its application to the spanish case. transportation research procedia. volume 12, 2016, pages 314-324. [5] laboratory of transport technics. czech technical university in prague, faculty of transportation sciences, department of transport telematics (16120), konviktská 20, prague 1, 110 00 czech republic, http://dsfd.fd.cvut.cz, 10.2.2017. [6] faller gmbh, gütenbach / schwarzwald. amtsgericht freiburg hrb610917, http://www.faller.de, 7.10.2016. [7] čsn 73 6101 projektování silnic a dálnic. praha: český normalizační institut, 2004. [8] m. a.-p.-taylor. intelligent transport systems. university of south australia, adelaide, http://www.emeraldinsight.com/doi/abs/10.1108/ 9781615832460-031, 10.2.2017. [9] conceptualizing smart city with dimensions of technology, people, and institutions. taewoo nam & theresa a. pardo, center for technology in government, university at albany, state university of new york, u.s., https: //www.ctg.albany.edu/publications/journals/dgo_ 2011_smartcity/dgo_2011_smartcity.pdf, 10.2.2017. 9 http://dsfd.fd.cvut.cz http://www.faller.de http://www.emeraldinsight.com/doi/abs/10.1108/9781615832460-031 http://www.emeraldinsight.com/doi/abs/10.1108/9781615832460-031 https://www.ctg.albany.edu/publications/journals/dgo_2011_smartcity/dgo_2011_smartcity.pdf https://www.ctg.albany.edu/publications/journals/dgo_2011_smartcity/dgo_2011_smartcity.pdf https://www.ctg.albany.edu/publications/journals/dgo_2011_smartcity/dgo_2011_smartcity.pdf acta polytechnica ctu proceedings 12:5–9, 2017 1 introduction 2 system overview 2.1 technical description of the model 2.1.1 physical layer of the model infrastructure 2.2 traffic control 2.3 model operation and control 3 further perspective 4 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0026 acta polytechnica ctu proceedings 5:26–28, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app traffic management system in terms of data exchange dušan kamenický czech technical university in prague, faculty of transportation sciences, department of transport telematics, konviktská 20, prague, czech republic correspondence: kamendus@fd.cvut.cz abstract. prediction of train running and it support of conflict resolution decision for an efficient use of the existing railway infrastructure is needed. to meet these requirements standardized interfaces between infrastructure managers and railway undertakings and infrastructure description are indispensable. keywords: technical specification for interoperability, infrastructure description, traffic planning, dispatching, train controlling, telematics applications, railml, railtopomodel. 1. introduction with the increasing demand for freight and passenger transport railway subjects aims on increasing the capacity by a reduction of delays and improved traffic fluidity. to meet these requirements modern traffic management systems are implemented. they are based on a prediction of train running and automatic conflict resolution. standardized interfaces shall ensure the needs of the railway undertakings, infrastructure managements and customers of different european countries in traffic planning, dispatching and train control. directive 2008/57/ec of the european parliament and of the council of 17 june 2008 on the interoperability of the rail system within the community defines in annex ii infrastructure, energy, control-command and signaling and rolling stock as structural subsystems of the rail system, and traffic operation and management, telematics applications for passenger and freight services as a functional subsystem of the rail system. specifications are drafted by the european railway agency and adopted in a decision by the european commission, to ensure the interoperability of trans-european rail system. the paper describes data exchange with focus on the operational management. 2. architecture of management process railway management process can be divided into three basic layers: traffic planning, dispatching and train controlling. traffic planning can be characterized as conceptual planning, strategic network development, service planning and infrastructure planning in long term, timetabling in short term. timetabling is a process concerning transport demand by customer and path requests and allocation by railway undertakings and infrastructure managers [1]. traffic planning is influenced by transport demands, infrastructure parameters, rolling stocks parameters and path fees. traffic planning shall define rules, like priorities of the train and competence between railway undertakings and infrastructure managers. the aim of dispatching is at close approximation to scheduled state, from which railway system has been deflected by external influences. it support is necessary in dispatcher decision processes [2]. quality and timeliness of information is necessary condition the dispatcher could correctly decide the intervention, which leads to fulfillment of the planned timetable. infrastructure manager dispatcher is responsible for changing train sequences to minimalize deviations from the timetable. railway undertaking dispatcher is responsible for skipping a commercial stop, breaking a connection, providing rolling stock and staff. major incidents forcing those decisions like rolling stock or infrastructural failures with a following adaptation of the timetable require a close communication between infrastructure manager and railway undertaking. train controlling is process train route setting, shunting route managing and setting, infrastructure elements controlling to ensure railway safety. routine processes can be performed by automatic train route setting system. automatic train control system, which is designed to eliminate human error, can be supplemented by automatic train operation system or driver advisory system. a train movement is adjusted to achieve defined points of infrastructure in defined time slot and therefore to reduce energy consumption and increase the capacity at a time. 3. standardized interfaces technical parameters of infrastructure elements, like line layout, track parameters, switches and crossings, 26 http://dx.doi.org/10.14311/app.2016.5.0026 http://ojs.cvut.cz/ojs/index.php/app vol. 5/2016 traffic management system in terms of data exchange platforms etc. are defined by 1299/2014/eu "technical specifications for relating to the infrastructure subsystem of the rail system in the european union" [3]. with exception of the requirements for infrastructure register, infrastructure description, methodology of representation or the storage of infrastructure data are not defined. infrastructure register is defined by 2014/880/eu "specifications of the register of railway infrastructure" [4]. this specification concerns data about the infrastructure structural subsystem the energy structural subsystem, and the trackside control-command and signaling subsystem. railway network shall be subdivided into sections of line and operation points. section of line means the part of line between adjacent operational points and may consist of several tracks. operational point means any location for train service operations, where train services may begin and end or change route and where passenger or freight services may be provided [5]. the structural subsystems specific requirements that interoperable railway lines and rolling stock must meet and many of these requirements must be stored in the infrastructure registers and the register of rolling stock. comparing those registers should make it clear which lines accept which rail vehicles. telematics applications for passenger and freight services subsystem equipment are defined by 2006/62/es "technical specification for interoperability relating to the telematics applications for freight subsystem of the trans-european conventional rail system" (tsi taf) [6] and 2011/454/es "technical specification for interoperability relating to the subsystem telematics applications for passenger services of the transeuropean rail system" (tsi tap) [7]. the specifications relating to telematics applications define architecture of information system and interfaces among subjects: infrastructure manager, railway undertaking and customer. processes and data exchanges to allocating of train path and monitoring train movement are defined. fundamental processes are: • train path request – path departure point, path departure time, path destination point, command and control system including on-board radio equipment, train weight and length, braking system and braking performance, maximum speed, rid numbers relating to any dangerous goods, information concerning exceptional gauging etc., • train composition message – for the preparation of the train, railway undertaking must have access to the infrastructure restriction notice, to technical wagon data (rolling stock reference database), to the dangerous goods reference file and to current updated information status on the wagons (the wagon and intermodal unit operation database), • train ready – message must be sent to indicating, that train is ready for departure, • train running information and train running forecast. interface between telematics applications and subsystem "control-command and signaling" is not specified, although interlocking equipment need data on the parameters of the trains. interface between telematics applications and subsystem "infrastructure" is given with the train path data definition and via the infrastructure restriction notice database. interface between telematics applications and subsystem "rolling stocks" is only given via the rolling stock reference database. the procedures enabling a coherent operation of various structural subsystems during both normal and degraded operation are defined by 2015/995/eu "technical specification for interoperability relating to the operational and traffic management subsystem of the rail system in the european union" [8]. from the viewpoint of the data exchange, relevant is requirement of route book and timetables. the development of the route book is the responsibility of ru and should be prepared in the language of the railway undertaking. however, the format of the route book is not specified. each train must be identified by a train running number. the train running number is given by the infrastructure manager when allocating a train path. the train running number format is defined in commission decision 2012/88/eu. interfaces among different railway it applications weren’t specified. the railml.org initiative was founded in 2002 in order to create an interface to enable heterogeneous railway applications to communicate with each other [9]. the result has been the development of the railway markup language railmlwhich delivers a universally applicable data exchange format. the railml standard has been developed by infrastructure managers, railway undertakings, software and consulting firms and academic institutions from number of countries. the railml specification contains subschemas for four main areas: infrastructure, timetable, rolling stock and interlocking. railml can be seen as a direct use case of the railtopomodel. the railtopomodel is a logical object model to standardize the representation of railway infrastructurerelated data [10]. on the other side, railtopomodel will become international railway standard (irs) in spring 2016, compiled by uic, the largest railway organization worldwide. sncf réseau has worked since 2011 to develop its global model for railways business objects, ariane model, based on the same principles as railtopomodel. ariane model now covers the description of all functional objects and properties of the network, at track and lines levels, including topology, referencing and topography, routes description and signaling. on 2016-2017 catenary and power supply network, and finalization of rolling stock will be added. 27 dušan kamenický acta polytechnica ctu proceedings 4. coclusion technical specification for interoperability defines technical requirements of infrastructure, energy and rolling stocks parameters and control-command and signaling system to enable interoperable vehicle driving on interoperable infrastructure. necessary data exchange relating to path allocation and operation and management processes relating to ensure safety are defined in function subsystems. however, data exchange relating to dispatching is not standardized. infrastructure description, specified in infrastructure register, is not sufficient for prediction of train running. data exchange relating to train parameters, like traction characteristics, and timetable requirements is not specified. infrastructure managers and railway undertakings interface ensuring dynamic data exchange relating to breaking a connection is not defined in technical specification for interoperability. railml seems to become effective tool for data exchange supporting prediction of train running and conflict resolution decision among others. references [1] birgit jaekel, thomas albrecht. operational railway management as part of an integrated railway management process, 2014. euro žel, žilina. [2] d. kamenický. optimalizační algoritmy pro systémy řízení a zabezpečení železniční dopravy, studie k disertační práci, 2015. čvut v praze, fakulta dopravní, praha. [3] 1299/2014/eu ”technical specifications for relating to the infrastructure subsystem of the rail system in the european union”. [4] 2014/880/eu ”specifications of the register of railway infrastructure”. [5] j. barnet. rail infrastructure informational description, master thesis, 2011. ctu in prague, faculty of transportation sciences, prague. [6] 2006/62/es ”technical specification for interoperability relating to the telematics applications for freight subsystem of the trans-european conventional rail system”. [7] 2011/454/es ”technical specification for interoperability relating to the subsystem telematics applications for passenger services of the trans-european rail system”. [8] 2015/995/eu ”technical specification for interoperability relating to the operational and traffic management subsystem of the rail system in the european union”. [9] railml website [online]. [2016-01-25], http://www.railml.org/en/. [10] railtopomodel website [online]. [2016-01-25], http://www.railtopomodel.org/en/. 28 http://www.railml.org/en/ http://www.railtopomodel.org/en/ acta polytechnica ctu proceedings 5:26–28, 2016 1 introduction 2 architecture of management process 3 standardized interfaces 4 coclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0012 acta polytechnica ctu proceedings 7:12–17, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app microscopic features of cement paste modified by fine perlite vladimír hrbeka, ∗, veronika koudelkováb, pavel padevěta, petr šašekb a czech technical university, thakurova 6, prague 6, czech republic b institute of theoretical and applied mechanics as cr, v.v.i., prosecka 76, prague 9, czech republic ∗ corresponding author: vladimir.hrbek@fsv.cvut.cz abstract. the use of waste material and replacement of binder element in cementitious composites is in focus of material development. perlite in the construction industry is usually used in form of lightweight aggregate enhancing the insulating performance of concrete. this paper focuses on integration of fine perlite into the cement matrix and possible replacement of the cement binder in the composition of the material. the macromechanical performance of the modified paste is tested on specimens with 5, 10, 15 and 20 % fine perlite substitution and pure cement sample. to distinguish the effect of the perlite on the microstructural level, pure cement material and specimen containing 10 % of fine perlite are investigated by the electron microscopy. furthermore, the mechanical properties of individual phases are examined and compared on same samples by instrumented indentation. the presented results enabled estimation of fine perlite impact on the macro and microscopic performance of the material. keywords: perlite, sem, nanoindentation, mechanical properties. 1. introduction reduction of energy consumption and co2 emissions production is the important challenge in the concrete industry at the present time, thus the focus on the utilization of different waste material partly replacing cement binder plays the key role in the field of research [1]. among supplementary materials perlite represents an expanded natural material produced through loosing water during heating of source material hydrous volcanic glass. most frequently is perlite with its porous structure used as a filler in the lightweight concretes. moreover, it also improves concrete insulation properties and fire resistance. processing of the crude natural volcanic material and the production of perlite lead into creation of perlite fines currently considered as the waste material [2]. the quality of concrete containing fine perlite is possible to asses from micro as well as from macro aspects. the most important phenomenon observable on the microscale is pozzolanic activity of perlite fines since creation of chemical bond between perlite particles and surrounding binder is crucial in developing proper mechanical properties of the cementitious mixture. the second important parameter is the total volume of fine perlite in cement which influences the mechanical behavior of cement composite. yu et. al [3] determined the most effective perlite content equal to 15 % since it significantly improved the compressive strength of concrete (about 16 mpa higher than in the case of concrete without perlite fines addition). similar results obtained rózycka and pichór [4] during testing of perlite waste addition effect on the properties of autoclaved aerated concrete. the compressive strength didn´t significantly changed up to 30 % of volume addition of perlite waste particles. on the contrary oktay et. al [5] determined that the highest compressive strength showed the concrete specimens with no addition of expanded perlite. the results mentioned above [3–5] confirms the importance of chemical availability of perlite particles resulting in creation of interface between perlite and surrounding binder. hence, this paper is aimed at the investigation of macroscopic and microscopic mechanical properties of cementitious paste containing different volume (5, 10, 15 and 20 %) of crushed fine perlite particles. the results of macroscopic testing determined samples for microscopical evaluation, where specimen with 10 % substitution of perlite is compared to pure cement paste. the microstructure is observed implementing the scanning electron microscope (sem), which enables defining different chemical phases and their interfaces in the material. the micromechanical properties of individual phases are determined from histograms of the results provided by instrumented indentation. 2. materials and methods the macroscopic specimens were prepared from ordinary portland cement cem i 42.5r mixture with water to cement ratio 0.35. the composition according to declaration of fine perlite producer is summarized in table 1. the specimens were placed in 40 × 40 × 160 mm forms, cast out 24 hours after mixing and stored until 28 days of total age in water 12 http://dx.doi.org/10.14311/app.2017.7.0012 http://ojs.cvut.cz/ojs/index.php/app vol. 7/2017 to prevent the carbonation. the sets of specimens are denoted as modcem x.xx, where x.xx stands for the fine perlite percentage substitution and specimen modcem 0.00 served as reference sample (pure cement paste). the residues from macroscopic testing were embedded in epoxy resin and used for microscopic investigation. the samples were sectioned by diamond cut-off wheel and the surface grinded and polished using silica-carbon papers (grid roughness p1200, p2400 and p4000 according to european p-grade system) and the 3 µm diamond suspension. chemical content granulometry sio2 min 66 % > 0.2 mm max 10 % al2o3 max 18 % > 0.1 mm min 50 % fe2o3 max 3 % < 0.1 mm max 50 % cao + mgo max 6 % na2o + k2o max 8 % table 1. composition of used fine perlite. 2.1. macromechanical testing to determine macromechanical properties i.e. bending strength and compressive strength, 6 macroscopic specimens (40×40×160 mm dimensions) were tested from each type of cement mixture. tensile stresscapacity was determined from load-controlled threepoint bending test with support spacing of 100 mm (l) from the equation 1, where ft is maximum applied force causing bending moment, b and h are dimensions of sample cross-section. σt = 3 2 ftl bh2 (1) compressive strength of samples was determined by load-controlled compression test with pressure pad area of 40×40 mm and calculated according the equation 2, where fc is maximum applied compressive force and a is the pressure pad area (40 × 40 mm). σc = fc a (2) figure 1. macroscopic tests setting – three-point bending (left), compression (right). 2.2. sem sem investigation was performed in mira ii lmu (tescan corp., brno) on polished specimens coated with thin layer of carbon necessary to ensure proper conductivity of the surface. working distance of the microscope was set closely to 15 mm and accelerating voltage was set to 15 kv to provide good signal. the micrographs were acquired at different magnifications. each specimen was investigated at first at the magnification 600× to gain good overview and than at higher magnifications (above 1200×) for better study of different phases in detail. sem micrographs acquired with back scattered electron detector (bse) provide information about distribution and chemical composition of different phases since back scattering coefficient strongly depends on the atomic number z. 2.3. nanoindentation to determine micromechanical properties, instrumented grid indentation using nanoindenter ti 750 serie (hysitron inc.) was performed on the polished specimens with high surface quality. in this technique mechanical properties are calculated from the unloading part of the force-displacement dependency diagram. one of the most important parameter which can be directly calculated is hardness h defined as pressure the maximum force (pmax) under the contact area of the tip (ac). h = pmax ac (3) the reduced (effective) elastic modulus (er) of measured volume of the material follows from the relationship between the unloading stiffness (s) and contact area with respect to the probe geometry (β) [6]. er = √ π 2β s √ ac (4) a grid indentation of 10 × 10 pattern with equally spaced indents (30 µm separation) was used in order to cover sufficient area containing all material phases. the load-controlled protocol of indentation consisted of loading holding unloading parts lasting 5 25 5 seconds respectively and reaching maximum applied force of 2.5 mn. to obtain proper statistical set of measured data, four grid indentations were performed on different places of each sample. mechanical properties of individual phases in the heterogeneous cement paste can be obtained from the histograms with equally spaced bins of the examined property by statistical deconvolution [7, 8]. in this paper, we assess the properties of individual phases based on difference between histograms of indentation results of modified cement paste (modcem 0.10) and referential specimen (modcem 0.00). 13 v. hrbek, v. koudelková, p. padevět, p. šašek acta polytechnica ctu proceedings 3. results 3.1. three-point bending, compression the results of macromechanical testing (table 2) correspond with results presented by oktay et. al [5], i.e. referential pure cement samples reach the highest values of compressive (81.89 ± 9.68 mpa) and tensilestress capacity (1.50 ± 0.83 mpa). mod/ref value present deviation of specimen mechanical property from referential sample. at this preliminary stage of research, perlite modification of cement paste with similar macromechanical performance was in focus. according to table 2, specimens containing 10% of fine perlite substitution (modcem 0.10) exhibit lowest decrease of both compressive and tensile-stress capacity (about 8% and 4% respectively). this led to selection of modcem 10 sample for microscopic investigation of the material. compressive strength specimen σc [mpa] denotion mean stat.dev. mod/ref modcem 0.00 81.89 9.68 1.00 modcem 0.05 64.01 13.28 0.78 modcem 0.10 76.04 10.84 0.93 modcem 0.15 64.69 5.42 0.79 modcem 0.20 70.80 11.32 0.87 tensile strength specimen σt [mpa] denotion mean stat.dev. mod/ref modcem 0.00 1.50 0.83 1.00 modcem 0.05 1.36 0.63 0.91 modcem 0.10 1.45 0.52 0.97 modcem 0.15 1.57 0.34 1.05 modcem 0.20 1.56 0.30 1.04 table 2. macromechanical test results. 3.2. sem investigation the sem equipped with energy dispersive x-ray detector (edx) enables to distinguish different phases in the studied cementitious materials (modcem 0.00, modcem 0.10). based on the bse micrographs and epma (electron probe microanalysis using edx) was possible to determine five phases clinker, csh gels, portlandite (calcium hydroxide ch), pores and perlite particles (see in the figure 2). clinkers in the gray-scaled bse micrographs represent the most intensive particles because of their dense structure and high iron and calcium content. around clinker particles is possible to determine darker well defined area called high density calcium-silica-hydrate gel (hd-csh) as a result of progressive slow hydration of non-hydrated clinker cores. the hd-csh gel represent the first one modification of csh, the second one is low density csh gel (ld-csh) with more opened porous structure. figure 2. bse micrograph showing microstructure of cement paste with perlite addition. the size of perlite was variable from approximately 200 µm of undamaged whole particles to perlite fines i.e. internal lamella of perlite with various size. the interface between undamaged perlite particles and surrounding cement matrix was sharp and didn’t show in most cases any chemical reaction detectable at microscale in the sem (figure 4). nevertheless, according to literature perlite fines are considered as good pozzolanic material [3]. in our investigation chemical reaction between crushed internal lamella and cement was observed in the specimen modcem 0.10. the evident structural disintegration coupled with decreasing content of si in affected perlite lamella is demonstrated in the figure 3. figure 4. bse micrograph showing microstructure of cement paste with perlite addition. 14 vol. 7/2017 figure 3. sem – eds of integrated perlite fines – si content decrease and detail of lamella integration. the decreasing tendency of si content can be also presented on epma with edx analysis results of perlite lamella and cement matrix boundary. figure show positions of testing and table describes weight percentage containment (wpc) of individual elements in each position of interest. position 4 with si containment 10.55% of weight correspond to csh phases and position 8 (si wpc 16.31%) to pure perlite [9]. figure 5. epma – edx of perlite lamella and cement. pos 8 7 6 5 4 ca 9.80 27.61 23.49 23.25 23.91 si 16.31 15.38 13.96 11.39 10.55 al 3.11 2.88 3.34 2.27 4.22 mg 0.05 0.06 0.23 0.19 0.21 k 4.09 1.75 4.29 4.65 5.28 na 0.76 1.26 1.12 0.80 0.73 s 0.30 0.40 0.36 0.71 0.42 fe 1.17 0.51 0.86 0.82 1.22 o 64.42 50.15 55.85 55.91 53.47 table 3. epma – edx chemical elements analysis (weight percentage containment in %). 3.3. instrumented indentation the previously described phases of the material (ld csh, hd csh, ch and clinker) and their mechanical properties (reduced modulus) can be observed in the indentation histogram. the ld csh forms a significant peak close to 30 gpa, which is consistent with previous research (21.7 ± 2.2 gpa [7, 8, 10]). measured indentation modulus of the hd csh phase is close to 40 gpa (occurrence of peaks in range 35 to 50 gpa) and the ch reaches up to 60 gpa, according to the figure 6. more accurately, using statistical deconvolution, reduced modulus estimation reach 27.34 ± 2.56 for ld csh phase, 39.29 ± 4.11 gpa for hd csh phase and 61.54 ± 4.30 gpa for ch phase. compare to phase reduced modulus measured on samples with w/c ratio 0.5 (29.4±2.4 gpa), hd csh phase reaches peak of indentation modulus naturally higher due to lower w/c ratio. the clinker phase is not in focus of this study due to its high stiffness and interference with other “low stiffness” phases, which misrepresent the result of indentation (collected data above 90 gpa are therefore excluded). figure 6. modcem 0.00 – reduced modulus histogram. figure7 represents the indentation data measured on specimen modcem 0.10. significant peaks of the cementitious matrix are slightly shifted to lower values due to the interaction of fine perlite residues to the composite structure and formation of “low stiffness" phase. the formation is not yet known to authors and will be subjected to further investigation. the 15 v. hrbek, v. koudelková, p. padevět, p. šašek acta polytechnica ctu proceedings individual phases of the cement form the peaks around 25 gpa for ld csh phase, 33 gpa for hd csh and 49 gpa in case of ch phase. the young’s modulus of expanded perlite is assumed to be equal 20 gpa [11, 12], which correspond to the peak found in the histogram. there is very significant peak formed around 10 gpa, which the authors have not been able to identify yet. the lower indentation modulus of all phases in modified cement paste also explains the decrease of macroscopic performance of the composite. figure 7. modcem 0.10 – reduced modulus histogram. 4. conclusions cementitious composite with substitution of fine perlite was investigated on macro and microscopical level. the macro mechanical testing of specimens showed slight decrease of material strength for all cases of cement replacement. the lowest decrease was observed on modcem 0.10, which was thus tested on microscopical level. the investigation of microstructure with sem proved the interaction of fine perlite lamella with cement matrix and integration of whole perlite particles. the instrumented indentation enabled to identify impact of cement paste modification due to comparison of referential data and reduced modulus of phases in specimen modcem 0.10. more detailed interaction of fine perlite and cement composite will be further researched. list of symbols σt bending strength [mpa] ft maximum applied force causing bending moment [n] b specimen cross section width [m] h specimen cross section height [m] σc compressive strength [mpa] fc maximum applied compressive force [n] a pressure pad area [m−2] l spacing of supports [m] h hardness [gpa] pmax maximum force [pa] ac contact area under the tip [nm−2] er reduced elastic modulus [gpa] β probe geometry constant [–] s stiffness [nm−1] acknowledgements the research has been supported by czech science foundation (project no. p105/12/g059). references [1] paris j.m., roessler j.g., ferraro c.c., deford h.d., townsend t.g., a review of waste products utilized as supplement to portland cement in concrete. j clean product 121:1 – 18, 2016. doi: 10.1016/j.jclepro.2016.02.013 [2] kotwica l., pichór w., nocún-wczelik w., study of pozzolanic action of ground waste expanded perlite by means of thermal methods. j therm anal calor 123:607 – 613, 2016. doi: 10.1007/s10973-015-4910-8 [3] yu l.h., ou h., lee l.l., investigation on pozzolanic effect of perlite powder in concrete. cem con res 33:73 – 76, 2003. doi: 10.1016/s0008-8846(02)00924-9 [4] rózycka a., pichór w., effect of perlite waste addition on the properties of autoclaved aerated concrete. const built mat 120:65 – 71, 2016. doi: 10.1016/j.conbuildmat.2016.05.019 [5] oktay h., yumrutas r., akpolat a., mechanical and thermophysical properties of lightweight aggregate concretes. const built mat 96:217 – 225, 2015. doi: 10.1016/j.conbuildmat.2015.08.015 [6] oliver w.c., pharr g.m., measurement of hardness and elastic modulus by instrumented indentation: advances in understanding and refinements to methodology. j mat res 19:3 – 20, 2004. doi: 10.1557/jmr.2004.0002 [7] constantinides g., chandran k.r., ulm f.j., vliet k.v., grid indentation analysis of composite microstructure and mechanics: principles and validation. mat sci eng a 430:189 – 202, 2006. doi: 10.1016/j.msea.2006.05.125 [8] němeček j., vondřejc j., králík v., micromechanical analysis of heterogeneous structural materials. cem con com 36:85 – 92, 2013. doi: 10.1016/j.cemconcomp.2012.06.015 [9] kabra s., katara s., rani a., characterization and study of turkish perlite. inter j innovative res sci, eng and techno 2:4319 – 4326, 2013. issn: 2319-8753 [10] velez k., maximilien s., damidot d., fantozzi g., sorrentino f., instrumented determination by nanoindentation of elastic modulus and hardness of pure constituents of portland cement clinker. cem con res 31:555 – 561, 2001. doi: 10.1016/s0008-8846(00)00505-6 [11] sengul o.,azizi s.,karaosmanoglu f.,ali m., effect of expanded perlite on the mechanical properties and thermal conductivity of lightweight concrete. energy build 43:671 – 676, 2011. doi: 10.1016/j.enbuild.2010.11.008 [12] abidi s., joliff y.,favotto c., impact of perlite, vermiculite and cement on the young modulus of a plaster composite material: experimental, analytical and numerical approaches. compos. b 92:281 – 36, 2016. doi: 10.1016/j.compositesb.2016.02.034 16 acta polytechnica ctu proceedings 7:12–16, 2017 1 introduction 2 materials and methods 2.1 macromechanical testing 2.2 sem 2.3 nanoindentation 3 results 3.1 three-point bending, compression 3.2 sem investigation 3.3 instrumented indentation 4 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0058 acta polytechnica ctu proceedings 7:58–63, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app experimental, numerical and analytical investigations of wind-induced net pressures for industrial buildings with envelope porosities vanessa saubke∗, rüdiger höffer ruhr-universität bochum, universitätsstraße 150, 44801 bochum, germany ∗ corresponding author: vanessa.saubke@rub.de abstract. the magnitude and the spatial distribution of wind-induced net pressures (external and internal) on buildings are frequently discussed among research communities and construction industries. this paper deals with this topic based on a case study about an industrial building in denmark, which was damaged due to the wind impact during a storm when a large part of the roof covering was blown off. in order to detect the reason for the damage the wind-induced loads were studied by i) wind tunnel experiments on the external pressures due to different wind directions, ii) analytical investigations of internal pressure due to envelope porosities and planned openings and iii) numerical analyses for the internal and the external pressure. the reynolds averaged navier-stokes (rans) method is employed to build a numerical model. the experimental, analytical and numerical results are compared with the indicated characteristic loads from the eurocode [1]. keywords: net pressure, internal pressure, industrial building, eurocode. 1. introduction the storm "christian" caused a lot of significant damages in northern europe in october 2013, including also failure of supporting structures of buildings. especially in denmark high wind velocities were measured. according to the danish meteorological service local wind peaks of 53.5 m/s were listed, which were the highest values ever measured in denmark. the wind velocity distribution for the maximum values for this day is given in figure 1 (b). the diagram in figure 1 (c) shows the data of the two measurement stations nearby the building which is considered here. the peaks of the wind velocity at 1 p.m. and at 3 p.m. are clearly visible. figure 1. (a) cardinal points (google earth [2], modified), (b) distribution of the peak wind velocity in denmark during hurricane "christian" in october 2013 (danmark meteorologiske institut [3], modified) and (c) measurements of the wind velocities at stations near to the building. at 3.15 p.m. the industrial building was damaged due to the wind impact. a large part of the roof structure breaks off and leads to an extended hole in the roof construction (see fig. 3 (a)). the debris were blown over the building and damaged also the leeward building (see fig. 2). the wind direction was west-southwest (for the cardinal points see fig. 1 (a)). figure 2. pictures from a security camera taken during the storm in 2013, the roof failure is visible. the damage started at the edge of the roof where the foil peeled off. this led to a deformation of the parapet wall (see fig. 3 (d)). approximately 7.5 % of the roof construction was completely blown off, additional 6 % of the roof foil was torn off and the trapezoidal sheets were folded and buckled (see fig. 3 (b)). in total more than 10 % of the roof construction was destroyed. futhermore this caused a damage at the sprinkler system and the leaking water destroyed a big part of the stored goods. the impact of the wind load was huge so that also the girders of the supporting roof construction were deformed (see fig. 3 (c)). 58 http://dx.doi.org/10.14311/app.2017.7.0058 http://ojs.cvut.cz/ojs/index.php/app vol. 7/2017 exp., num. and anal. investigations of wind-induced net pressures figure 3. qualitative description of typical damages: (a) partically opened roof, (b) buckled sheets, (c) bent girder and (d) deformed parapet. the considered building is located in the east part of denmark. it is situated at the west of a complex of three storage buildings of the same type. each building is about 135 m long, 55 m wide and 38 m tall. the shelfs for the stored goods are part of the supporting steel construction which consists of a regular framework of stiffening frames (see fig. 4). this main structure is covered by a sandwich-element facade. figure 4. example of a warehouse building under construction (google maps [2]). the roof consists of supporting girders, trapezoidal sheets, mineral insulating panels and a roof foil. the sheets are connected to the supporting girders by screws and the foil is fixed by spikes. for the sandwichelement facade a connection type with unsealed joints was used. after lange and rädel [4] this specific joint construction can lead to a leakage of the building walls. for the consideration of this leakage an effective joint was estimated as 3.5 mm (for investigations on the leakage of joint see kuhnhenne [5]). with this information the general porosity φ, the relation of the area of the openings to the area of the surfaces, was calculated: φ = ∑ a′∑ a = 0.0026 = 0.26%. (1) at the southern facade the gates for the delivery of the goods are located behind an antecedent flat hall. if the gates are not closed the total area of openings is 61.43 m2. 2. materials and methods 2.1. wind loads after 1991-1-4 after the eurocode external and internal pressures due to wind impact are taken into account [1]. a superposition of both leads to the net pressure. the definition of the signs of this values are given in figure 5. figure 5. three typical cases for the internal pressure (cook [6], modified). the code defines the wind load we,i as multiplication of the peak velocity pressure and a pressure coefficient: we = qp · cpe and wi = qp · cpi. (2) the peak velocity pressure qp is influenced by different factors like the terrain category, the building height and the wind zone. the external pressure coefficient cpe is influenced by the shape of a building and the internal pressure coefficient cpi by the distribution of the openings over the four faces of the building. in the following investigations pressure coefficients are considered and compared. for this purpose the external pressure coefficients for the roof of this building are determined in table 2. figure 6. distribution of the pressure coefficients for a flat roof with (a) wind direction on the short side and (b) wind direction on the long side. regarding the flow around a rectangular cube of a certain height (see hucho [7]) the code estimation of 59 vanessa saubke, rüdiger höffer acta polytechnica ctu proceedings the distribution of the pressure coefficients is given in figure 6. the distribution depends on the dimensions of the building and has to be estimated for every building and every wind direction individually due to the regulations. the cpe,1 values should be used for loaded areas of less than 1 m2 and the cpe,10 values for loaded areas of more than 10 m2 with the possibility of a logarithmic interpolatin for intermediate values. the internal pressure can be calculated according to two different regulations: (1.) for homogeneously distributed openings the pressure coefficient can be determined from a diagram by calculating the opening ratio µ, which is the relation from the area of the openings in the suction region divided by the area of all openings: µ = ∑ area of openings with cpe ≤ 0∑ area of all openings (3) (2.) a face with an area of openings that is more than twice the area of the sum of the openings in the remaining faces is considerd as a dominant face. • the internal pressure coefficient can be calculated with the following equation: cpi = 0.75 · cpe. (4) • if the area is more than three times the remaining openings then it is calculated with the following equation: cpi = 0.9 · cpe. (5) the considered building has firstly a general porosity because of the leakage due to the facade-elements. secondly it has large openings for the delivery of the stored goods at the short face where the open flat building is located. it is not known yet if openings were open or closed during the storm. therefore, three different situations for two wind directions are considered. these different cases are explained in table 1. the results for the internal pressure coefficients for these six cases are listed in the table 3. case windward face porosity delivery gates 1 x 2 short x x 3 x 4 x 5 long x x 6 x table 1. description of the six load cases. 2.2. experimental tests wind tunnel tests were performed at the boundary layer wind tunnel at ruhr-universität bochum to build up a reference data-base of wind-induced loads on the structure for specifically chosen spots of the building envelope, especially on the parts damaged during the storm event. the focus of the measurements was set on the roof construction which failed (see fig. 7 (a)). the effect of the parapet wall was neglected due to the small size in relation to the building. this leads to results on the safe side. figure 7. (a) positions of the pressure measurement points and (b) the principal set-up of the model. the tests were performed in a geometrical scale of 1:300 for different wind directions from southwest to northwest. the building was investigated as a (i) single standing, isolated building and (ii) in a group arrangement with the neighboring buildings (see fig. 7 (b) for the set-up). the evaluated cpe values represent the 78 % fractiles of the analyzed measurement data. the pressures are measured over a point-like averaging area which represents an area smaller than 1 m2 in full scale. the minimum values from the flow direction on the long face and of all measured directions for every zone are listed in table 2. figure 8. (a) the minimal cpe values of the flow direction west (on long face) and (b) of all measured flow directions. 2.3. numerical simulations for the numerical investigations a cfd method applying finite volumes and the rans method with the standard k-�-model is used. the applied inlet velocity profile was introduced by richards and hoxey [8] in 1993 and is commonly used for this method. effects of building interferences were neglected by running all simulations with a single building. figure 9. (a) idealization of the joints and (b) idealization of the openings. 60 vol. 7/2017 exp., num. and anal. investigations of wind-induced net pressures because a simulation of all explicit openings and joints would lead to an extremely high number of volume elements the openings were idealized. the general porosity of the joints was considered with four explicit idealized joints distributed over the height (see fig. 9, left). the openings of the delivery gates were also idealized as one large opening with the same area which is located in the center of gravity of all gates (see fig. 9, right). figure 10. numerical results for the internal pressure coefficients cpi: (a) case 1, (b) case 2, (c) case 3, (d) case 4, (e) case 5 and (f) case 6. figure 11. numerical results of the external pressure coefficients of the roof for (a) flow on the short side and (b) flow on the long side. geometrical models and meshes were generated for all six cases (see table 1) with approximately three million volume elements per case. the distribution of the external pressure coefficients is pictured in figure 11. the maximum values for every zone are listed in table 2. the internal pressure coefficients for the six different cases are plotted in figure 10 and numbered in table 3. 2.4. analytical calculations an analytical investigation is performed for the estimation of the internal pressure. as cook [6] 1990 already described, the internal pressure pi depends on the external pressure pe at the openings a′. it is defined as the balance between the inflow and outflow of the building (see also fig. 5 (c)): inflow∑ ( a′ (pe − pi) 1 2 ) = outflow∑ ( a′ (pi − pe) 1 2 ) . (6) the equation cannot be solved directly but iteratively with an estimated value for the internal pressure. the left-hand and right-hand side of equation 6 are not fixed during the iteration process, because an outflow can become an inflow and vice versa. figure 12. distribution of the external pressure coefficients from a wind tunnel test for (a) the wind direction on the short face and (b) the wind direction on the long face (tpu aerodynamic database [9], modified). to calculate the internal pressure analytically a distribution of the external pressure is needed. because the performed wind tunnel tests are putting the focus on the damaged area of the roof construction, complete roof and wall distributions of external pressures have not been measured and are missing for analytical calculations of internal pressures. therefore mean values of external pressure coefficients from wind tunnel tests of a building with similar proportional dimensions from the tpu aerodynamic database [9] were used (see fig. 12). because the building from the database was smaller than the original one, the area of the openings was scaled down whereas the opening ratio µ stayed unchanged. in this scale the distributions of the pressure coefficients were evaluated and the area of the different pressure coefficients calculated. in order to maintain the comparability the same idealized opening of the delivery gates like those for the numerical investigations in section 2.3 were used. for the joints a general porosity of 0.26 % was considered. the internal pressure was calculated for all six cases as given in table 1 after cook’s method, [6]. the results are listed in table 3. 61 vanessa saubke, rüdiger höffer acta polytechnica ctu proceedings 3. results the results of the wind tunnel investigations correspond relatively good to the required values of the eurocode, for example the eurocode requires a value of cpe = 2.2 for the zone f and the wind tunnel results leads to the highest value of cpe = 2.14 considering all investigated directions. the results are compared to the cpe,1 values although the area of the measurement points in the experiments correspond to an area of smaller 1 m2 in full scale. the normative values of zone f and g suit to the minimum values of the testing results for all directions while the normative value of zone h fits better to the averaged value. but as one can see in figure 8 the values of this zone have a certain variation, they decrease in direction to the leeward side of the building. therefore, it is obvious that a design for the maximum value would not be economic, but one should be aware of the possibility that higher values can occur. zone f g h eurocode1) -2.2 -1.8 -1.2 windtunnel test (long face) -1.64 -1.34 -1.43 (-1.192)) windtunnel test (all directions) -2.14 -1.84 -1.63 (-1.292)) numerical simulation (case 1-3) -1.19 -1.11 -1.01 numercial simulation (case 4-6) -1.18 -1.13 -1.03 table 2. comparison of the external pressure coefficients cpe for the roof construction. 1)cpe,1 values in order to maintain comparability 2)averaged value of zone h the cpe values of the wind tunnel tests for the wind on the long face can be compared to the numerical results of cases 4 to 6. it should also be taken into account that the rans method does not lead to peak values and deliver average value only. the distribution of the external pressure coefficient shows a satisfactory similarity to the experimentally determined distribution from the windtunnel tests. but rans underestimates the external pressure in the separation area because the turbulence is modelled mathematically and local instationary vortices are neglected. the values of the numerical results are between 72 % to 84 % of the values of the wind tunnel test. the investigations of the internal pressure coefficients lead to different results for the six cases. the magnitude of the values are similar, in most of the cases, like case 1, 3 and 6, the eurocode overestimates the values and is on the safe side, but in case 5 it underestimates the internal pressure in comparison to the analytical solution. case 1 2 3 4 5 6 windward face short long porosity x x x x delivery gates x x x x eurocode -0.30 0.34 0.64 -0.03 -0.33 -0.83 numerical simulation -0.21 -0.12 0.52 -0.24 -0.31 -0.63 analytical calculation -0.20 0.34 0.59 -0.23 -0.48 -0.62 table 3. comparison of the internal pressure coefficients cpi for the six considered load cases. in case 2 the numerical analysis lead to a complete different value than the results of the analytical solution, which seems to fit perfectly to the value of the eurocode. in a further investigation, e.g. with the les-method, the numerical results should be checked, but it is sure, that this is connected to the position of the opening. if the position of the opening in the analytical calculation is changed to a higher position in the middle of the face (see fig. 13), the value increases from cpi = 34 to cpi = 0.45. figure 13. changed position of the opening at the windward face (tpu aerodynamic database [9], modified). also in some additional numerical investigations on a 1m x 1m x 1m cube, where the position of the openings over the height was changed, the same behaviour was determined (see fig. 14). this shows that the internal pressure reacts quite sensitive to the position of the openings, especially for openings in the windward side. figure 14. investigation on a 1m x 1m x 1m cube due to the position of the opening: (a) opening at the bottom, (b) opening in the middle and (c) opening at the top. in case 4 the eurocode diagram leads to a value of nearly zero (with µ = 0.63 in fig. 15) in contrast to the numerical and analytical solution, which leads to good corresponding results of -0.24 and -0.23. in 62 vol. 7/2017 exp., num. and anal. investigations of wind-induced net pressures the old german code near to zero values were avoided. in the range of the diagram where the value leads to zero one has to consider two values (0.2 and -0.3) on the safe side (see the red square in fig. 15 (b)). figure 15. diagram for the cpi values: (a) eurocode [1] and (b) the previous code din 1055 [10]. 4. conclusions the results show, that the regulation of the internal pressure for a dominant face is conservative for dimensioning a building (see regulation (2.) in section 2.1). the regulation for homogeneously distributed openings however seems to falsify the internal pressure (see regulation (1.) in section 2.1). as well for very homogeneously distributed openings, like a general porosity, the regulation (1.) does not work properly, like in case 4. also in case 5 the value is underestimated in comparison to the analytical value. the normative din en 1991-1-4 [1] offers the opportunity to use the fixed values of +0.2 and -0.3, if it is not possible or not justifiable to estimate a µ. these values seems to represent an underestimation of the effects of the internal pressures as one can understand from the results of the investigations for cpi ranging from +0.59 to -0.63. this apparent load underestimation through the quoted simplified code regulation does not necessarily lead to damages of the structure as it can be assumed to be based on experiences and sound practice requirements which are included into code regulations. however, such configuration can be insecure in connection with a certain systematical underestimation of external pressures after specific design regulation for trapezoidal sheets. for example, the german din 18807-3 [11] considers the higher pressures at the edge of the roof (zone f and g) just for the design of the screw connections; for the design of the sheets the high suction in these zones is neglected, which leads to an underestimation of the external pressure at the edge of the roof. together with the simplified eurocode regulation as described above this can result into a safety relevant lack of net design pressures. the research on the topic of the internal pressure should be continued to validate the results. to perform some additional wind tunnel tests on buildings with different proportions in combination with numerical simulations would be useful. for the numerical simulations some further methods should be used to improve the results. the aim would be to find a stable model that leads to good results for the different proportions so that it can be used as a tool to check and improve the regulation of the code. list of symbols a area of the faces [m2] a′ area of the openings [m2] cpe external pressure coefficient cpi internal pressure coefficient pe external pressure [n/mm2] pi internal pressure [n/mm2] we external pressure (eurocode) [n/mm2] wi internal pressure (eurocode) [n/mm2] µ opening ratio φ general porosity references [1] e. c. for standardization. en 1991-1-4 eurocode 1: actions on structures part 1-4: general actions wind actions. 2010. [2] google maps. panoramio, 2007. [3] danmark meteorologiske institut (dmi). orkan satte danmarksrekord i vind se kortene, 2013. [4] j. lange, f. rädel. die fugendichtigkeit von sandwichelementkonstruktionen wasserund luftdichtigkeit in längsfugen und fensteranschlussfugen. in festschrift zum 60. geburtstag von gerhard hanswille, heft 20. institut für konstruktiven ingenieurbau, bergische universität wuppertal, 2011. [5] m. kuhnhenne. energetische qualität von gebäudehüllen in stahl-sandwichbauweise. dissertation, rheinisch-westfälische technische hochschule aachen, fakultät für bauingenieurwesen, 2009. [6] n. j. cook. the designer’s guide to wind loading of building structures part 2: static structures. butterworths, 1990. [7] w.-h. hucho. aerodynamik der stumpfen körper. vieweg+teubner, 2nd edn., 2011. [8] p. richards, r. hoxey. appropriate boundary conditions for computational wind engineering models using the k-� turbulence model. journal of wind engineering and industrial aerodynamics 46-47:145–153, 1993. [9] t. p. u. aerodynamic database. database of isolated low-rise building without eaves. http://www.wind.arch.t-kougei.ac.jp/info_ center/windpressure/lowrise/mainpage.html. [10] din 1055-4 action on structures part 4: wind loads. din deutsches institut für normung e.v., 2005. [11] din 18807-3 trapezoidal sheeting in buildings; steel trapezoidal sheeting; strength analysis, structural design. din deutsches institut für normung e.v., 1987. 63 http://www.wind.arch.t-kougei.ac.jp/info_center/windpressure/lowrise/mainpage.html http://www.wind.arch.t-kougei.ac.jp/info_center/windpressure/lowrise/mainpage.html acta polytechnica ctu proceedings 7:58–63, 2017 1 introduction 2 materials and methods 2.1 wind loads after 1991-1-4 2.2 experimental tests 2.3 numerical simulations 2.4 analytical calculations 3 results 4 conclusions list of symbols references 170 acta polytechnica ctu proceedings 1(1): 170–174, 2014 170 doi: 10.14311/app.2014.01.0170 black hole results from xmm-newton norbert schartel1 1xmm-newton science operations centre, esa, villafranca del castillo, apartado 78, e-28691 villanueva de la cañada madrid, spain corresponding author: norbert.schartel@sciops.esa.int abstract xmm-newton is one of the most successful science missions of the european space agency. since 2003 every year about 300 articles are published in refereed journals making directly use of xmm-newton data. all xmm-newton calls for observing proposals are highly oversubscribed by factors of six and more. in the following some scientific highlights of xmm-newton observations of black holes are summarized. keywords: x-ray agn black hole. 1 introduction xmm-newton ([1]) is the second cornerstone of european space agency’s (esa) horizon 2000 science programme, providing an observatory-class x-ray facility. the spacecraft was launched by an ariane 5 on 10 december 1999. the observatory provides simultaneous non-dispersive spectroscopic imaging and timing (european photon imaging camera; epic, [2] and [3]), medium resolution dispersive spectroscopy (reflection grating spectrometer; rgs, [4]) and optical/uv imaging, spectroscopy and timing from a co-aligned telescope (optical monitor; om, [5]). the three x-ray mirrors ([6]) in combination with the cameras of epic offer a large effective area over the energy range from 300 ev to 12 kev, up to 2500 cm2 at 1.5 kev and ∼1800 cm2 at 5 kev. the scientific potential of the effective area may be illustrated by the first observation of an evolving dust-scattered x-ray halo around a gamma ray burst ([7]). each of the two modules of the rgs cover the energy range from ∼0.4 kev to 2.2 kev with an effective area of 60 cm2 at 15 å. 2 scientific highlights scientific highlights resulting from the first decade of xmm-newton and chandra observations can be found in [8]. in the following i list a number of highlights from xmm-newton observations of black holes. in this paper i have focused on some of the most exciting discoveries in this field, which also received wide publicity in public relations announcements by esa. 2.1 galactic black holes and ultraluminous x-ray sources globular clusters (gc), containing thousands of stars packed within tens of light years, were considered as a possible breeding ground for black holes. a rival hypothesis suggests that black holes are ejected through close star encounters and consequently gcs are devoid of black holes. xmm-newton observations of ngc 4472 allowed the first detection of a black hole in a gc ([9]) excluding the latter hypothesis. ultraluminous x-ray sources (ulx) were proposed to harbour intermediate-mass black holes, which provides the link between stellar mass black holes and supermassive black holes (smbh) in the centres of galaxies. xmm-newton and chandra observations of cxom31 j004253.1+411422 in andromeda allowed to connect this ulx to low mass x-ray binaries ([10], implying accretion onto a stellar-mass black hole in the eddington regime. a second ulx found in andromeda, xmmu j004243.6+412519, allowed to observe the emission of the accretion disk in x-rays together with the emission form its jets in radio ([11]). the experimental key for both observations was the low absorbing column density towards andromeda whereas absorption is a major obstacle of ulx observations in our own galaxy. [12] found an intermediate-mass black hole in ngc 1313. the x-ray spectra of the ulx can be described with a power law plus an accretion disk (kt ∼= 150 ev) implying a mass of ∼= 103m�. a intermediatemass black hole with m > 500 m� could be associated with an ulx in eso 243-49 based on luminosity variations observed by xmm-newton ([13]). 170 http://dx.doi.org/10.14311/app.2014.01.0170 black hole results from xmm-newton 2.2 the strong gravitational field currently, x-ray observations are the only way to observe the strong gravitational field in the direct vicinity of black holes and neutron stars ([14]). special and general relativistic effects distort the spectra of particles orbiting black holes depending on the orbital parameters and the black hole’s spin. theoretical discussion can be found in [15], [16], [17] [18] or more recently in [19] and [20]. experimentally the iron kα line is most studied as there are no lines of abundant elements nearby. early examples are the xmm-newton observations of the galactic black hole xte j1650-500 in outburst ([21]) and the active galactic nuclei (agn) mgc-6-30-15 in low state ([22]). both spectra are explained by a fast spinning black hole and the extraction and dissipation of rotational energy from it. the simultaneous xmm-newton and nustar observation of ngc 1365 revealed that reflection from an ionized disk readily explains the spectra taken by both satellites ([23]). the xmm-newton spectra of the galactic black hole, gx 339-4 in outburst ([24]) is an example of a black holes with almost maximum spin. the observation of 1h 0706-495 is unique as it shows not only the iron kα, but also the iron lα line ([25]). in addition the light cure shows the expected characteristic variability of reflection from an ionized disk (compare also [26], [27], [28], [29]). 2.3 active galactic nuclei (agn) in low states following [30] x-ray spectra of agns are composed of a power-law continuum emitted above a black hole plus reflection from an ionized disk. during the low state the continuum emission region moves nearer to the black hole and gravitational bending affects its light path. observationally, during low states the continuum emission appears suppressed whereas the reflected emission appears constant or even enhanced, compare also [31]. the most intensive studied agn in low state with xmm-newton is pg 2112+059, where an additional layer of ionized material was used to favour the reflection interpretation versus alternative scenarios ([32], [33] and [34]). [35] used variability considerations to discriminate the reflection interpretation versus an absorption scenario for pg 0844+349 in an x-ray weak state. and [36] could demonstrate for the low state observation of 1h0707-495 reflected emission within one gravitational radius of the event horizon of the black hole. 2.4 aspects of variability near supermassive black holes (smbhs) whereas quasi-periodic-oscillations (qpo) are well established in x-ray binaries for almost 30 years, qpos remained elusive in agns. xmm-newton measured a ∼1 hour qpo for re j1034+396 ([37]). [38] found evidence for orbital motion of material close to the central black hole of mrk 766. a xmm-newton observation allowed [39] to observe a co-rotating flare at a distance of only 3.5 to 8 schwarzschild radii to the smbh of ngc 3516. 2.5 energy budget, winds and outflows [40] established for the fist time simultaneous spectral energy distributions for the majority of the [41] reverberation mapped sample of agn based on xmmnewton epic and om measurements. [42] used xmmnewton observations to show that radio-galaxies produce sufficient mechanical energy to unbind a significant fraction of the intra-group medium, an effect which is negligible in massive clusters of galaxies. combining high resolution rgs spectra with sensitive light-curves of epic, [43] demonstrated an accretion-disk origin for the two warm absorber winds in ngc 4051. 1h 0707495 shows a mildly relativistic, highly ionized outflow which changed its velocity from about 0.11c to 0.17c between 2008 january and 2010 september ([44]). ultrafast outflows are present in >35% of radio-quiet agn observed with xmm-newton, providing a significant contribution to the agn cosmological feedback ([45], [46], [47]). 2.6 flares and tidal disruption events [48] observed several peaks in the power density spectrum of the x-ray light curve of the smbh in the galactic centre during which period a bright x-ray flare was detected ([49]). theoretical studies revealed a previously unknown topological structure inherent to black holes with high spin: in a small region near the event horizon of the spinning black hole the orbital velocity decreases for decreasing orbital radius ([50]). this effect is now rightfully known as the aschenbach effect ([51]). [52] could identify a tidal disruption event based on rosat, chandra and xmm-newton observations. suzaku and xmm-newton observations taken shortly after the occurrence of the tidal disruption event swift j164449.3+573451 reveal a 200-second x-ray quasiperiodicity ([53]). this qpo might be explained with the forming of an accretion disc or precession of the jet. 171 norbert schartel 2.7 deep fields and cosmology a total of 1000 agn detections from a variety of rosat, xmm-newton and chandra surveys allowed [54] to obtain for the first time reliable space densities for low-luminosity (seyfert-type) x-ray sources at high cosmological redshifts. their evolutionary behaviour shows strong dependency on the x-ray luminosity and differs from the dependency found for high luminosity agns and quasars. xmm-newton allows the detection of quasars at highest redshift, e.g. sdss j104433012502 at z=5.80 ([55]). the spacecraft could even establish an x-ray spectrum of sdss j1030+0524 at z=6.30 ([56]). an ionized iron kα absorption edge in the x-ray spectrum of apm 08279+5255 allowed to obtain an, at the time of publication, highly interesting constrain on the age of the universe ([57]). 3 discussion and conclusions since 2003 every year about 300 articles are published in refereed journals making directly use of xmm-newton data. all xmm-newton calls for observing proposals are highly oversubscribed by factors of six and more. within esa’s mission extension scheme all missions are evaluated every 2 years and possibly extended by 4 years subjected to midterm confirmation. xmmnewton is funded up to end of 2016 subject to midterm confirmation and further extension discussion in 2014. currently, the xmm-newton mission is implementing four-reaction-wheel operation schemata, which will reduce fuel consumption significantly. the envisaged operation mode will allow technically operating the mission up to 2026. acknowledgement i thank the organizers of the workshop and especially franco giovannelli for giving him the opportunity to show highlights of xmm-newton black hole observations. and i thank the anonymous referee for many fruitful comments and suggestions. references [1] f. jansen et al. xmm-newton observatory. i. the spacecraft and operations. a a, 365:l1–l6, january 2001. 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[57] g. hasinger, n. schartel, and s. komossa. discovery of an ionized fe k edge in the z=3.91 broad absorption line quasar apm 08279+5255 with xmm-newton. apj, 573:l77–l80, july 2002. 174 http://dx.doi.org/10.1103/physrevd.71.024037 http://dx.doi.org/10.1126/science.1223940 introduction scientific highlights galactic black holes and ultraluminous x-ray sources the strong gravitational field active galactic nuclei (agn) in low states aspects of variability near supermassive black holes (smbhs) energy budget, winds and outflows flares and tidal disruption events deep fields and cosmology discussion and conclusions acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0099 acta polytechnica ctu proceedings 12:99–103, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app artificial neural network for models of human operator martin růžek czech institute of informatics, robotics and cybernetics (ciirc), czech technical university in prague, zikova 4, 166 36 prague 6, czech republic correspondence: martin.ruzek@cvut.cz abstract. this paper presents a new approach to mental functions modeling with the use of artificial neural networks. the artificial neural networks seems to be a promising method for the modeling of a human operator because the architecture of the ann is directly inspired by the biological neuron. on the other hand, the classical paradigms of artificial neural networks are not suitable because they simplify too much the real processes in biological neural network. the search for a compromise between the complexity of biological neural network and the practical feasibility of the artificial network led to a new learning algorithm. this algorithm is based on the classical multilayered neural network; however, the learning rule is different. the neurons are updating their parameters in a way that is similar to real biological processes. the basic idea is that the neurons are competing for resources and the criterion to decide which neuron will survive is the usefulness of the neuron to the whole neural network. the neuron is not using "teacher" or any kind of superior system, the neuron receives only the information that is present in the biological system. the learning process can be seen as searching of some equilibrium point that is equal to a state with maximal importance of the neuron for the neural network. this position can change if the environment changes. the name of this type of learning, the homeostatic artificial neural network, originates from this idea, as it is similar to the process of homeostasis known in any living cell. the simulation results suggest that this type of learning can be useful also in other tasks of artificial learning and recognition. keywords: neural network, artificial neuron, learning algorithm, mental model. 1. introduction for many practical applications it would be useful to have a model of a human operator, as for example in transport engineering where the human operator plays a principal role. unfortunately, the human factor causes also most of the accidents; therefore it would be helpful to understand the processes in the human brain. due to the complex nature of the human brain it is however very difficult to model and predict its behavior. he need to dispose of the model of a human operator, as the most complex part of many transport systems, stays behind this research. the transport and traffic simulations are on high level, but the principal component of any transport system – the human factor – is still quite unexplored. the main obstacles are the lack of knowledge about the mental processes and the differences between the typical artificial intelligence approaches and the biological neural networks. the artificial neural networks (ann) seems to be a promising method for these models because the architecture of the ann is directly inspired by the biological neuron and also because it is data driven method. the classical paradigms of artificial neural networks are however not suitable for direct use because they simplify too much the real processes in biological neural network, but it is possible to update the learning algorithm so that it corresponds more frankly to the biological reality. the neural networks have already been used in several projects aimed at mental models. one of them is the blue brain project[1, 2] with the goal to simulate the whole brain. this model should be so detailed that every cell is depicted on a molecular level. according to the authors, processes like consciousness, creativity, emotions or aggression will emerge. another research aimed at the models of human mind is the darpa synapse project with opposite attitude. the idea behind this research is to create a neural network with architecture similar to biological neural network, however with strongly simplified neurons[3]. the expectation is not to model mental processes, but to use the principles of biological neural network to improve the computer architecture. the way of information processing in the computer architecture is different from biological networks. in brain, the memory and computation are distributed; there is no central processor or memory. the expectation of synapse project is to understand the advantages of this architecture and to use them in formation of novel computer architecture. with respect to both above mentioned researches, the target of this paper is to find such an equilibrium point where the model of the neuron is still faithful enough so that it can model the strong processes and 99 http://dx.doi.org/10.14311/app.2017.12.0099 http://ojs.cvut.cz/ojs/index.php/app martin růžek acta polytechnica ctu proceedings yet be simple enough to be practically realizable. 2. new solution – homeostatical neural network the need for a compromise between the complexity of biological neural network and the practical feasibility of the artificial network led to a proposal of new learning algorithm. the idea is based on the classical multilayered neural network (mlp), the difference is in the learning process. the neurons are updating their parameters in a way that is similar to the real biological processes. the basic idea is that the neurons are in competition for resources and the survival criterion is the usefulness of the neuron to the whole neural network. the neurons are not using any "teacher" or other kind of superior system, they have the same information as the biological neuron. the learning process can be seen as searching of some equilibrium position which represents a state where the importance of the neuron for the neural network is maximal. this position can change in time if the environment changes. the name of this type of learning, the homeostatic artificial neural network, is derived from the similarity of this idea to the process of homeostasis that is known from biology. the proposed neural network is based on the idea of mcculloch-pitts neuron: y = φf n∑ i=0 xiwi (1) where f is the transfer function. the sigmoid transfer function was used: f(x) = 2 (1 + e−α∗x) − 1 (2) as the similarity to biological neuron was the basic requirement, the back propagation algorithm is not a solution because a higher structure (or teacher) is used to train the neuron. in the so called homeostatic neuron, the unit is using its proper forward connection to improve its function. the ’axon’ in this model has two functions – the first one is the transmission of the output to higher layer (as in back propagation), the second is the transmission of the utility information from higher layer to the lower one. the idea of this type of training is that the neuron improves its relative importance in the network, in other words, it is trying to maximize the part of its output signal that is accepted by other neurons. this idea corresponds to the biological reality because the information transmitted by the axon has the form of energy (and is inseparable from energy). therefore, the neuron knows which part of its output energy was accepted by other neurons. the process of learning can be described by the following algorithm: first, the neuron computes its output with its initial random weight. then the neurons in the higher layer set their weights according to their level of contentment with the reference neuron. in the next step, the neuron changes its weights accordingly. several possibilities of the weight change are described later in this paper. then the new output is computed. the neurons in the higher layer read the output and re-calculate their inputs weights. the reference neuron then decides which setting was better. this algorithm has several variants, all of them are using at least 2 successive values of the level of acceptance. this implies that the neuron must be equipped with a memory. the process of learning is on figure 1. from the point of view of the reference neuron the learning is described by the following algorithm: (1.) random initial weights (2.) output with initial weights for the first input (3.) neurons in higher layer (output neurons) compute the utility of the reference neuron and set their input weights accordingly. if they are satisfied with the output of the reference neuron, they increase their weights, otherwise they decrease them. there are several ways how to compute the utility, some of them are described in eq. 3 6. (4.) the reference neuron changes one (or more) of its input weights (5.) the reference neuron repeats the forward phase with the same data but with changed weight (6.) the output neurons compute the utility (as in step 3) (7.) the reference neuron evaluates the change in the step 4 – if it improves the utility, it will keep it, otherwise it will change the weights in the opposite direction (8.) neuron repeats steps 2 to 7 with all the connections and all the inputs. 3. criterions for optimization several methods can be used for the calculation the importance of the neuron. the basic difference among them is the number of output neurons for which the reference neuron is ’working’. the first extreme is a neuron that is optimizing its function for all output neurons. this neuron is ’reading’ all output weights without taking any particular weight into consideration. the opposite extreme is neuron that works only for one neuron in the higher layer; in other words is optimizing its function to improve its utility for this particular neuron. apart from these options, many other compromise criterions can be defined. 100 vol. 12/2017 artificial neural network for models of human operator figure 1. wide figure[4, 5]. 101 martin růžek acta polytechnica ctu proceedings this first idea corresponds to a neuron that is finding such a weight vector ∧ = w1,w2, ...,wn for which the sum of the absolute values of the output weights is maximal. this idea corresponds firmly to the biological reality because the neuron has only one axon and therefore it can only be aware of the total amount of the signal that is accepted by other neurons, not of the particular weights. in the case of the artificial neuron we also expect negative weights; therefore the neuron sums the absolute or square values. the utility q is: q = n∑ j=1 |woj | (3) respectively: q = n∑ j=1 (woj ) 2 (4) eq. 4 puts stress on great values and reduces the importance of the small ones. this may be in certain cases advantage for the learning but does not correspond to the biological reality. 4. searching the one neuron maximum – the second extreme the other type of training is based on the presumption that the neuron is increasing its importance to only one neuron in the higher layer, therefore it maximizes the function: q = max|woj |; j ∈ [0, 1, ...n] (5) the problem is that if max|wo| = 1, no further improvement is possible and the training stops. in a real situation, we expect networks with many neurons where this value will be reached very soon and then the learning stops. this is not the desired behavior; therefore in that case there should be used an additional condition that ensures the continuation of the training. the solution of this problem is to use a compromise that takes into consideration more than one output neuron but not all of them. this can be done by optimization of some given number of maximal output weights: u = max(|wo|) + max(wo −max(wo)) + ...; (6) where w = {w0, ...,wn−1}. 5. variants of homeostatical neural network apart from the above mentioned variant, several other options must be defined. first of all, the question of the weight range arises at least three options. the most biologically plausible is to limit the weights to interval between 0 and 1. that means that the weights cannot change the polarity of the signal and no gain of the output energy is allowed. this option is the most biologically plausible as the axon, synapse and dendrites are mostly working only as transmitters of the signal. on the other hand, some operation with the signal may be done also on the level of the axon dendrit transmission, for example the inhibitory weights can reduce the neuron potential. this leads to the second idea, to limit the weights to interval 〈−1; 1〉. the last option is not to limit the weights at all, that means the signal can be multiplied by any real number. this is the least biologically plausible method; on the other hand it will bring the highest computational power. the other question is how to choose the weights that should be updated. the first option that comes to mind is to update the weight that has the greatest influence on the result; that means the most sensitive weight. this will lead to highest increase of the utility of the neuron in the next step, however it may not be the best option from the global point of view. making the step always in the direction of the highest gradient may lead to falling into local extreme. also, to find the most sensitive weight consumes a lot of computational power. alternative option is to choose the updated weight randomly or in given order. these two options lead to similar results. it is also possible to imagine the combination of these algorithms, for example by using several steps to update of the most sensitive weights which may be followed by randomly chosen weights. because of the complex nature of the neural networks, it is not possible to say which method is the best for arbitrary data. 6. problems of the idea of homeostatical learning the described idea has two principal limitations. the first one is the learning of the highest layer. the idea of a learning algorithm is that neurons in certain layers are updating their weights according to the higher layer. this can be used for all layers except for the highest as it does not have any output neuron. the solution for practical simulation and testing is that the highest layer was not trained by the homeostatic learning algorithm, but by the back propagation. this is of course is an alteration of the original idea, but for the sake of the practical realization it is the easiest way how to program the homeostatic learning for the rest of the network. in the case of practical application, for example in robot, this problem will be solvable in a natural way because the network will be part of a closed loop. the second problem is the delay question. the reference neuron is updating its input weights accord102 vol. 12/2017 artificial neural network for models of human operator ing to its output weights, but in each neural cell the output calculation takes some time., meanwhile the inputs are changing this means that the forward information will not ’meet’ the information about the utility in the same time. this problem has two solutions. the first is to set the dynamics of inputs to a lower level, so that the speed of information processing in the whole system is significantly higher than the changes of the input signals. the other possibility is to equip the neurons with a memory that stores the previous inputs so that it is possible to recall them when the information about the utility reaches the neuron. 7. conclusions in this paper, the idea of a new learning algorithm for neural networks is presented. this algorithm can cope with some disadvantages that the classical neural networks paradigms have. the proposed algorithm has several variants; from the point of view of the optimization criterion it is optimization for all input weights vs. optimization for only one input weights, and many compromising solutions. from the point of view of the weight update, the possible variants are the update of the weight with the highest sensitivity to the change, of randomly chosen weight of consequently chosen weight. from the point of view of the weight range, it is possible to define weights as positive numbers smaller than 1, as either positive or a negative number in absolute value smaller than 1, or as real numbers. every variant has its pros and cons. due to the high number of variants which is even multiplied by the amount of data (for neural networks as for data driven method it is important which data is fed into the network. the same network can work well with one data and wrongly with another) it is difficult to decide which variants is the most promising. larger testing is needed to understand the quality of the methods. despite the fact that the initial tests showed that the signal prediction task is better fulfilled by back propagation algorithm, the homeostatic neural network seems as a promising method for modeling on mental processes. of course it is not possible to create a model of the complete consciousness, but it is possible to concentrate on some specific region. transport engineering brings many possible applications as the human factor is the most important part in many transport systems. several processes may be modeled and predicted by the use of this network. typical example is a car driver, who is often making decisions with a lack of information and with the use of prior data. it is difficult to understand the factors that make for example the decision whether to cross the crossroads if the traffic light is orange and will become red in a short period of time. in such situations some drivers react differently even if the conditions are the same. the analytical way to explain such states is quite complex, but it is possible to collect enough data (either from real traffic or from a simulator) to predict this type of decisions. the neural network is a good tool to process this data. once the network is trained for certain tasks in a driver’s decision making processes, it can be used for the investigation of other decisions and behavioral predictions. this can help to explain and avoid danger situations caused by aggressiveness, fatigue, emotions and others. references [1] m. růžek, t. brandejský. model of biological ann based on homeostatic neurons. in 12th wseas international conference on neural networks (nn’11), isbn 978-960-474-292-9, athens 2011, pp. 66 69. [2] m. růžek. artificial neural network inspired by homeostasis in biological networks. in proceedings of 17th international conference on soft computing (mendel 2011), isbn 978-80-214-4302-0, pp. 232 235. [3] m. růžek. artificial neural networks for models of driver’s brain functions. in 20th anniversary of the faculty of transportation sciences, czech technical university in prague – selected papers, čvut v praze, 2013, isbn 978-80-01-05320-1, pp. 207 211. [4] m. růžek. modeling of eeg signal with homeostatic neural network. in nostradamus 2013: prediction, modeling and analysis of complex systems, ostrava 2013, isbn 978-3-319-00541-6 pp. 175 180. [5] m. růžek, t. brandejský. model of homeostatic artificial neuron. in neural networks, fuzzy systems & evolutionary computing, isbn 978-960-474-195-3, athens 2010, pp. 145 148. 103 acta polytechnica ctu proceedings 12:99–103, 2017 1 introduction 2 new solution – homeostatical neural network 3 criterions for optimization 4 searching the one neuron maximum – the second extreme 5 variants of homeostatical neural network 6 problems of the idea of homeostatical learning 7 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2015.1.0029 acta polytechnica ctu proceedings 2:29–33, 2015 © czech technical university in prague, 2015 available online at http://ojs.cvut.cz/ojs/index.php/app x-copter studio michal koutný∗, ondřej pilát, patrik černý, maroš kasinec charles university in prague, faculty of mathematics and physics, ke karlovu 3, prague, czech republic ∗ corresponding author: xm.koutny+pair@gmail.com abstract. we present a project that aggregates various existing robotic software and serves as a platform to conveniently control a quadrocopter, mainly for research or educational purposes. user interface runs in a browser and other components are also made with portability in mind. we provide a common interface that unifies different quadrocopter models and we implemented it for the parrot ar.drone 2.0. the platform is data oriented, i.e., it is based on dataflow between user objects. we implemented several such objects for: data recording and replaying, inertial and visual localization and following a given path. keywords: robotics, quadcopter, ide. 1. motivation despite the fact that cheap hardware (such as parrot ar.drone 2.0 [1] or ready kits [2]) is available, there are not many possibilities for application programmers to develop software for these robots without need to distinguish between individual models. our aim is to provide a platform for development of software for quadrocopters. the target users are ai programmers or students and expected tasks are general algorithms (basic example in figure 1) for quadrocopters. the result should work with any robot compatible with our software (section 4.3). the testing should be further simplified by running the application without physical access to a quadrocopter either by using a simulator or data previously captured during live flights. at(xcheckpoint.reachedcheckpoint) { xcheckpoint.checkpoint = nextcheckpoint(); },; // x = 3 m, y = 1 m, z = 1.5 m var cp0 = checkpoint.new(3, 1, 1.5); xcheckpoint.checkpoint = cp0; figure 1. sample script that flies through computed checkpoints. it assumes there is created a dataflow graph with the node xcheckpoint in it. 2. related work 2.1. middleware for robotics probably most popular middleware for robotics is the robot operating system (ros) [3]. it supports communication between objects using publish–subscribe mechanism. it is open source, mainly targeted on linux platforms. similar project is urbi sdk, developed by former private company gostai [4]. the current maintainer is aldebaran robotics [5], unfortunately the community around urbi sdk is much smaller and much less active in comparison with ros. reasons why we chose urbi sdk despite this fact are in the section 3.2. 2.2. parrot ar.drone 2 api part of our project is an api for parrot ar.drone 2. various other projects are dealing with this. there is the official sdk [6] with c++ api, controltower [7] that provide java interface, a ros package ardrone_autonomy [8], uobject for urbi sdk [9] or implementation of czech technical university [10]. because none of the aforementioned fit to our requirements for os portability, stability, functionality or documentation, we implemented our own (see the section 4.3). 2.3. ground control system there is official application for parrot ar.drone [11] intended for mobile phone users allowing manual control and displaying only limited data from sensors. the pc application controltower [7] allows controlling quadrocopter with specialized computer peripheries and has airplane-like gui. more complex application is qgroundcontrol [12] that cooperates with pixhawk project [13] that encompasses own hardware and uses visual localization. 3. used technology 3.1. architecture our system is divided into three components. first interacts directly with a user, second controls the robot and the last one connects the former two. the components are separate processes that communicate with each other over network, with intention to run components on different machines. 29 http://dx.doi.org/10.14311/app.2015.1.0029 http://ojs.cvut.cz/ojs/index.php/app m. koutný, o. pilát, p. černý, m. kasinec acta polytechnica ctu proceedings 3.1.1. client client software is used to create and launch user scripts, edit them, manually control the robot and visualize data (e.g., directly from quadrocopter’s sensors). despite the technology challenges the client is thin – running in a web browser.1 3.1.2. server the server component conveys communication between the client and the actual control machine (further onboard). its task is to control the access to the onboard and monitor quality of the connection between the client and the onboard. in the case the overall connection latency (client–server and server– onboard) exceeds preset limits, a warning message is shown to the user. if the connection is lost, onboard execution is correctly terminated and user is notified too. 3.1.3. onboard the onboard is the main executive component. the robot control and data processing run here because it is closest to the robot. the onboard is executing commands obtained from the client and sends back various data selected by user. the onboard component is supposed to run under normal operating system.2 our implementation exploits a pc that communicates with quadrocopter via wi-fi. we did not test the onboard component directly on a robot.3 3.2. urbi sdk urbi sdk is a c++ middleware for robotics, which we based the onboard component on. basically it provides support for communication between user objects (uobjects) and schedules user jobs. communication is possible via so called uvars, which are slots of uobjects. sender just writes to these slots and a receiver’s callback handles the change of uvar’s value. this allows both apparently asynchronous communication and really asynchronous when a thread pool is used to run the callbacks. further, uobjects can run in different processes and urbi sdk ensures transparent messaging via tcp or udp sockets. orchestration user scripts (written in urbiscript) can be executed by the urbi runtime. urbiscript is a prototype-based object-oriented language conceptually similar to javascript. it is possible to implement uobject functionality exclusively in the urbiscript. we chose urbi sdk because of its portability (linux and windows systems are supported) and the own 1google chrome is strongly recommended, though mozilla firefox will also get by (without visualization of video data). 2we support windows 7/8 and ubuntu 14.04 systems. 3the hardware of parrot ar drone 2.0 theoretically should be able to run our onboard with limited performance. scripting language and runtime. considered alternative was robot operating system. 3.3. nodejs nodejs is a server-side javascript engine. recently, it became quite popular among developers of interactive web applications and various modules [14] exist that extend core functionality. it suited our needs for the server component. 3.4. html5 thanks to the standardization efforts many features that were earlier common only for desktop applications or via third party plug-ins (flash, java applets, native plug-ins) are now implemented directly in the browser, generally referred to as html5. to make client as multi-platform as possible, we decided to implement the client for the browser using aforementioned html5 technologies. most importantly, we use the web socket api [15] for sending data back to the client and <video> tag to display streamed video. 3.5. c++ the executive parts of the onboard component are written in c++ (which is consequence of using urbi sdk). we utilize features of the c++11 standard, mainly for threading and memory management. 4. features 4.1. dataflow graph we chose dataflow driven approach to describe the user’s program. it consists of units of operation which we call nodes. each node has at least one input or at least one output. any operations are either results of changes on the node inputs or are launched by an external event hidden in a node (e.g., timer expiration, a socket received data). urbi sdk itself encourages dataflow control by its communication paradigm. we extended the original mechanism to ensure syntactic and semantic compatibility between nodes. dataflow bears following advantages for the end user. 4.1.1. implicit parallelism nodes are implicit synchronized on outputs-inputs connections, thus avoiding excessive user effort. this also ensures certain level of scalability on multiple cpus machines. 4.1.2. separating abstraction levels user can concentrate on high level objectives only, implementation details are hidden in particular nodes. on the other hand, a creator of a node is limited by a particular node interface. 30 vol. 2/2015 x-copter studio 4.1.3. visual development each node defines an interface, which makes it possible to check what connections are possible (not only syntactically but also semantically with flat system of semantic types). in the end, user can assemble various dataflow graphs interactively (figure 2). 4.1.4. in-time data inspection selected outputs of the nodes can be connected with a special node that resends the data to the gui where the data are visualized on the fly (figure 2). 4.2. scripting only the dataflow graph of nodes is not enough for controlling the quadrocopter. thus, a user can write scripts that are sent and executed on onboard. in such scripts it’s possible to handle events occurring within the dataflow graph or manually affect the dataflow. the script execution can be paused and resumed (that is a feature of urbiscript, for instance useful for intentionally infinite loops, see figure 3 for urbiscript console). 4.3. quadrocopters unified api the x-copter studio supplies a unified application programming interface (denoted as xci) for data acquisition and control of quadrocopters. it is ready for common sensors on quadrocopters and it can be easily extended. we implemented the xci for one of the most popular quadrocopters – parrot ar.drone 2.0 and also a simulated quadrocopter in the v-rep simulator [16]. our parrot implementation exploits asynchronous socket events handling, which allows us to detect connection failures and attempt to restore the connection with the quadrocopter. the implementation further supports full configuration of the parrot ar.drone 2.0 and reading data from all available sensors. we use parrot for its price, robustness and popularity among research groups. the cost for this is extensibility – neither the hardware nor the software on parrot can be expanded. we have to deal with unreliable communication over wi-fi and connection failures. physical dimensions of the ar.drone 2 with the hull are 53 cm × 52 cm and it weights 420 g including battery. it can fly for about 15 minutes with the more powerful battery version (1500 mah). the ar.drone 2 provides a video stream in high definition (720p) with 30 fps from front camera aimed forward or qvga stream with 60 fps from bottom camera aimed to the ground, three-axes gyroscope, three-axes accelerometer, three-axes magnetometer, pressure sensor and ultrasound altimeter. the quadrocopter sends raw and adjusted measurement from gyroscope, accelerometer, ultrasound sensor and pressure sensor with frequency of 200 hz. all data from these sensors are accessible for user in the dataflow. quadrocopter firmware uses aforementioned sensors to maintain tilt, rotation and vertical velocity of the quadrocopter according to an external fly control command. this control command u = (φ, θ, ψ, ż) ∈ [−1, 1]4 consists of quadrocopter’s roll φ, pitch θ, yaw rotation speed ψ and vertical velocity ż. it should be sent every 30 ms for smooth movements. 4.4. data recording and replaying thanks to the structured dataflow architecture, any potentially interesting data can be captured and stored in a file. we use text format which is portable and allows convenient processing with data processing tools. data that are inherently binary aren’t supported, however, special case – image data are stored with compression to a separate video file. recorded data are stored together with timestamps, so it’s possible to “replay” them and test or debug system’s response with real timing. 4.5. localization in unknown environment we use quadrocopter’s onboard sensors to give the end user information about the absolute position of the aircraft. the localization is a node within the dataflow graph, thus it runs exclusively on the onboard component. 4.5.1. inertial localization our implementation expects accelerometers with aggregated outputs in a form of horizontal velocity, altimeter and drone pose sensors (i.e., accelerometers for roll and pitch angles and gyroscope and/or magnetometer for yaw angle). sensor data are filtered with extended kalman filter (ekf) where we engage physical model based on a similar project [17]. 4.5.2. visual localization to enhance absolute position estimate, we are processing the video stream from the onboard camera. we use monocular slam framework ptam [18] that estimates camera pose and position for each frame and also builds a map of observed feature points. first, the map is initialized from a pair of images where point correspondence is based on optical flow and the initial scale is estimated thanks to inertial localization. after the initialization, the scale estimate is continuously refined using method described in [17]. data both from inertial and visual localization are merged together in the extended kalman filter (see figure 4) in separate update steps. the result from ptam (absolute position and quadrocopter’s pose) is used to update the ekf, taking into account a delay of visual localization and then recalculating prediction of the current state. 31 m. koutný, o. pilát, p. černý, m. kasinec acta polytechnica ctu proceedings figure 2. x-copter studio gui. top navbar indicated connection quality, upper half is filled with widgets that display data from the onboard, lower half accommodates the editor with repository of nodes on the left, and connected nodes on the right. at the bottom there are (left to right): dataflow control buttons, scripting console control buttons and manual flight control buttons. the scripting console is hidden. figure 3. urbiscript console. on the top – script editor, middle – output of executed script, bottom left – script execution control (surrounded with other controls of x-copter studio). 4.6. following given path path is represented as a list of checkpoints that drone strictly follows. a checkpoint is a composite of world coordinates and a tangential vector.4 4this vector specifies direction of quadrocopter’s heading at the given checkpoint. this is intended to be used when checkpoints should be followed as a spline curve. checkpoint attainment quadrocopter achieves a checkpoint when it is in a sphere around the checkpoint with 10 cm diameter (according to the localization) and immediately continues to the next checkpoint in the queue. flight to the checkpoint is controlled by 200 hz pid regulator.5 regulator uses data from the extended kalman filter in order to eliminate data delay and predict quadrocopter’s current position in the world. output from the regulator is sent directly to the quadrocopter. 4.7. runtime configuration any configuration (parameters, calibration constants, settings, etc.) is stored in human readable and editable text files and loaded during runtime, thus avoiding necessity of recompilation to apply changes. the same configuration can be accessed both from c++ and urbiscript api. 5. conclusion we hope the x-copter studio will become used (at least) for educational purposes since it provides sufficiently high-level interface, for instance to prototype planning algorithms. 5the frequency is induced by the dataflow, in this case it is refreshed frequency of data from quadrocopter. 32 vol. 2/2015 x-copter studio -0.5 0 0.5 1 1.5 2 2.5 3 -1 -0.5 0 0.5 1 1.5 2 2.5 3 x [m] -0.5 0 0.5 1 1.5 2 2.5 3 -1.5 -1 -0.5 0 0.5 1 1.5 2 2.5 x [m] y [m ] heading path heading path figure 4. example of a path tracked by the ekf. a square 2 m × 2 m was set up and quadrocopter was navigated (with pid controller) to its vertices using inertial localization only. supposed position is shown on the left plot. as you can see, plot on the right shows more realistic pose and position thanks to visual localization (applied to the original data). alas, we don’t have any reference data for absolute comparison. if it is proved as useful and practical, it is possible to extend the software for other types of robots, not only quadrocopters. references [1] parrot. ar.drone 2.0. parrot new wi-fi quadricopter. [2015-02-08], http://ardrone2.parrot.com/. [2] bitcraze. the crazyflie nano quadcopter. [2015-02-08], http://www.bitcraze.se/crazyflie/. [3] open source robotics foundation. ros.org. [2015-02-08], http://www.ros.org/. [4] gostai. urbi. [2015-02-08], http: //www.gostai.com/products/jazz/urbi/index.html. [5] aldebaran. aldebaran/urbi. [2015-02-08], https://github.com/aldebaran/urbi. [6] ardrone open api platform. [2015-02-08], https://projects.ardrone.org/. [7] javadrone – ar.drone java api – google project hosting. [2015-02-08], https://code.google.com/p/javadrone/. [8] m. monajjemi. ardrone_autonomy – ros wiki. [201502-08], http://wiki.ros.org/ardrone_autonomy. [9] urbiforge projects/urbi 4 ar drone. [2014-08-28], http://www.urbiforge.org/index.php/projects/ urbi4ardrone. [10] t. krajník, v. vonásek, d. fišer, j. faigl. ar-drone as a platform for robotic research and education. in research and education in robotics eurobot 2011, pp. 172–186. springer science + business media, 2011. doi:10.1007/978-3-642-21975-7_16. [11] ar.freeflight. [2015-02-08], http://ardrone2.parrot.com/#freeflight. [12] l. meier. qgroundcontrol gcs. [2015-02-08], http://www.qgroundcontrol.org/. [13] l. meier, p. tanskanen, l. heng, et al. pixhawk: a micro aerial vehicle design for autonomous flight using onboard computer vision. auton robot 33(1-2):21–39, 2012. doi:10.1007/s10514-012-9281-4. [14] node packaged modules. [2015-02-08], https://www.npmjs.org/. [15] a. m. i. fette. rfc 6455 – the websocket protocol. [2015-02-08], http://tools.ietf.org/html/rfc6455/. [16] c. r. gmbh. coppelia robotics v-rep: create. compose. simulate. any robot. [2015-02-08], http://www.coppeliarobotics.com/. [17] j. engel, j. sturm, d. cremers. camera-based navigation of a low-cost quadrocopter. in 2012 ieee/rsj international conference on intelligent robots and systems. ieee, 2012. doi:10.1109/iros.2012.6385458. [18] g. klein, d. murray. parallel tracking and mapping for small ar workspaces. in 2007 6th ieee and acm international symposium on mixed and augmented reality. ieee, 2007. doi:10.1109/ismar.2007.4538852. 33 http://ardrone2.parrot.com/ http://www.bitcraze.se/crazyflie/ http://www.ros.org/ http://www.gostai.com/products/jazz/urbi/index.html http://www.gostai.com/products/jazz/urbi/index.html https://github.com/aldebaran/urbi https://projects.ardrone.org/ https://code.google.com/p/javadrone/ http://wiki.ros.org/ardrone_autonomy http://www.urbiforge.org/index.php/projects/urbi4ardrone http://www.urbiforge.org/index.php/projects/urbi4ardrone http://dx.doi.org/10.1007/978-3-642-21975-7_16 http://ardrone2.parrot.com/#freeflight http://www.qgroundcontrol.org/ http://dx.doi.org/10.1007/s10514-012-9281-4 https://www.npmjs.org/ http://tools.ietf.org/html/rfc6455/ http://www.coppeliarobotics.com/ http://dx.doi.org/10.1109/iros.2012.6385458 http://dx.doi.org/10.1109/ismar.2007.4538852 acta polytechnica ctu proceedings 2:29–33, 2015 1 motivation 2 related work 2.1 middleware for robotics 2.2 parrot ar.drone 2 api 2.3 ground control system 3 used technology 3.1 architecture 3.1.1 client 3.1.2 server 3.1.3 onboard 3.2 urbi sdk 3.3 nodejs 3.4 html5 3.5 c++ 4 features 4.1 dataflow graph 4.1.1 implicit parallelism 4.1.2 separating abstraction levels 4.1.3 visual development 4.1.4 in-time data inspection 4.2 scripting 4.3 quadrocopters unified api 4.4 data recording and replaying 4.5 localization in unknown environment 4.5.1 inertial localization 4.5.2 visual localization 4.6 following given path 4.7 runtime configuration 5 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0068 acta polytechnica ctu proceedings 12:68–73, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app cross-study research on utility and validity of driving simulator for driver behavior analysis michal matowicki∗, ondřej přibyl czech technical university, faculty of transportation sciences, na florenci 25, prague 1, czech republic ∗ corresponding author: matowicki.m@gmail.com abstract. driving is one of the most ordinary and universal everyday tasks and, at the same time, one of the most complex and dangerous. it requires a full range of sensory, perceptual, cognitive, and motor functions, all of which can be affected by a wide range of stressors and experience levels. therefore, exploring of human behaviour while controlling a vehicle is a crucial task in improving traffic safety. experimental studies can always be conducted with on-road tests, however, using a simulator is safer and more cost-effective. the main goal of this paper is to demonstrate if and under what conditions could a driving simulator provide sufficient results required for a proper study of driver behavior. it discusses its limits and advantages. overall, the research reviewed in this paper indicates that simulator driving behaviour approximates (relative validity of speed and lateral position of vehicle on road), but does not exactly replicate (absolute validity), on-road driving behaviour. keywords: driving simulator, relative validity, absolute validity, cross-study analysis. 1. introduction with rapid growth of technological advancement, not only executive elements or traffic engineering like intelligent transport systems, is shifting toward more sophisticated and ecologically and economically valid solutions. another example is research on the transport subjects which is reaching for advanced tools like driving simulators. however, although supported in many government and scientific organizations [1] the usefulness and validity of results achieved in a driving simulator remain a very hot topic in transport engineering especially when driving behaviour and the psychology of the drivers are concerned. although this topic seems to be of significant importance (either supports of diminishing effects and results of driving simulator experiment), it seems to be often forgotten during the execution of the experiment, which may, at times, lead to meaningless conclusions. hence, it is very important to produce a literature review on the topic in search of significant data to utilize simulator results to make conclusions on real-world driver behaviour. this document includes a table (table 1) which summarizes the most important papers on the topic, as well as further analysis and description of chosen papers, all of which are divided into several categories depending on the kind of validity confirmed and parameters measured in the experiment. the table allows for quick familiarization with the most important research reviewed and introduces an in depth analysis of specific studies. chapter two consists of subsections that are dedicated to particular kinds of validity which can and should be of concern when setting up a driving simulator experiment. this paper produced a cross-study review and analysis of confirmed and proven validity measures for driving simulator experiment measures, and is concluded by summarizing recommendations and predictions regarding the future of driving simulators as tools for various studies on driving behaviour. 2. reviewed studies despite the importance that is placed on the simulator as a tool for investigating driver performance and behaviour, an extensive literature review reveals that only a very limited number of studies specifically evaluating the behavioural validity of simulators was conducted [3]. majority of the available literature on simulator validation has focused on measures of speed, lateral position, and braking responses (table 1). the following summary of literature research reviews literature on the given topic with a specific emphasis on the description of performance measures, the tasks and conditions, and first of all the evidence provided for behavioural validity of each simulator and the behavior of the subject "driving" it. 2.1. speed adjustment speed is one of the most studied measures of driver behaviour and, thus, one of most commonly analised with respect to simulator validity [2, 5–7, 9–11, 13, 15, 17]. overall, these studies rather consistently provide results confirming relative but not absolute validity of speed data in driving simulators. most explanatory ones were for example klee et al. [10] who examined the validity of a fixed-base simulator with respect to forward speed. 30 drivers drove the same road section on both driving simulator and in an instrumented vehicle. the speed values were measured at 16 points, while measures in 10 of them were similar in both observed environments. while authors do not conclude relative validity of the simulator and focus on lack of 68 http://dx.doi.org/10.14311/app.2017.12.0068 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 cross-study research of driving simulator authors achieved validity bella, 2008 [2] speed and complexity of menoeuvre mullen, charlton, devlin & bedard,2011 [3] relative validity of lateral position across different road types watts & quimby, 1979 [4] relativity of perceived hazard risk blaauw, 1982 [5] absolute speed validity & relative lateral displacement riemersma et al., 1990 [6] relative validity of speed reduction measures tornros, 1998 [7] relative validity of speed and lateral positioning wade & hammond, 1998 [8] relative validity of lane position reed & green, 1999 [9] absolute validity on speed control relative validity on effects of phone dialling klee et al., 1999 [10] relative validity of vehicle velocity bittner et al., 2002 [11] relative validity of speed in road curves blana & golias, 2002 [12] no relative validity confirmed godley et al., 2002 [13] relative validity of speed countermeasures lee, lee & cameron, 2003 [14] relative validity of visual attention-age relation bella, 2005 [15] relative validity of speed mcavoy 2007 [16] no validity for speed adjustment in work zones shinar & ronnen 2007 [17] relative validity of speed adjusting table 1. studies elaborating driving simulator validity. absolute validity, it can be observed that speeds in all 10 locations with similar results were 5-10 km/h slower for simulated driving. moreover, described speed difference provided the same results in each case which implies relative validity of simulator. another author providing evidence for the relative validity of the simulator maintained speed and its correlation with the on-road speed is bella [2, 15]. bella provides two studies in which the simulators were designed with identical road sections to those where real driving was performed by experiment participants. in real world experiments, speed measurement points were specified along the road instead of the instrumented vehicle driving record. the first experiment was conducted on road section with construction zone with speed limitation applied, whereas the second study described a highway in normal condition. a bilateral z-test for non-matched samples was performed in both data sets to estimate whether the recorded driving speeds were significantly different from those in the simulator. consistently with other researches bella conducted that simulator speeds were different than in the real world, although in [2] mean speed in simulator was higher than in real world while in [15] it was on the contrary. nevertheless, the differences in simulator speeds and in the field speeds for each point were not significantly different, demonstrating at minimum, interactive relative validity of static simulator tool for assessing speed. unconventional way of assessing speed as simulator validity indicator was presented by shinar & ronen [17] where two speed related measures were observed: speed estimation (i.e. driver perception of the current speed of vehicle) and production of speed (i.e. adjustment of the vehicle’s speed to reach predetermined speed). sixteen participants took part in an experiment where both on-road driving and simulator (with shielded speedometer) drives were performed. during each drive, participants were repeatedly asked to estimate the current speed of vehicle and produce seven different speeds 69 michal matowicki, ondřej přibyl acta polytechnica ctu proceedings ranging from 40 to 100 km/h. authors found that estimated speeds differed significantly in the simulator only for speeds of 40 and 100 km/h where the first speed was always estimated higher in simulator and latter one was always estimated lower. on the other hand, second investigated variable showed consistently higher produced speeds in simulator with a difference of approximately 25 km/h. conclusions of multiple analysed studies pointed out towards absolute validity of most of the estimated speeds and vast majority of the studies suggests relative validity for speed produced in the simulator. this seems to form firm foundations for performing scientific research with the use of driving simulators. 2.2. lateral position of vehicle second often validated measure in driving simulators is the lateral position of the vehicle on the simulated road. one of the most interesting studies in the matter of lateral positioning validity was performed by wade and hammond [8]. the authors measured positioning deviation as distance from the centreline of the road. in a tested sample of 26 participants, the mean deviation of the distance from centreline was statistically significantly larger in virtual driving when compared to the on-road conditions. authors determined that in the real-world it is clear that drivers travel more accurately near the centreline when compared to virtual driving experience. this led to the conclusion that absolute validity was not demonstrated in performed research. nevertheless, mullen, charlton, devlin and bedard [3] noticed that similar to results in speed researches mean deviations on different types of roads showed a familiar pattern when compared to the real environment (i.e. the road with the highest mean deviation in the real-world was also the road with the highest mean deviation in the virtual environment) which indicates relative validity. however, it is a noticeable fact that data were collected on straight road section with a relatively low speed limit of 35 mph (approx. 56 km/h). although blana & golias [12] tried to answer the question of speed influencing validity of lateral displacement measures, it remains unknown. in their research scientists conducted experiment in a simulator where distance between front wheel of the vehicle on the passenger side and white line determining the edge of the road were measured. a sample of 100 on-road drivers in vehicle instrumented with video cameras pointing toward road surface drove on elaborated highway section, while corresponding road was also driven in simulator by 100 participants. the results clearly pointed towards higher average distance measured in the real-world conditions when compared to simulator on both straight sections and curves. this indicates that drivers in the on-road conditions position their vehicles closer to the centreline of the road lane. this conclusion was further confirmed by smaller standard deviations (sd) for the measures recorded in real-world conditions. although the direction of displacement from the road edge was consistent for curved and straight road sections for all speeds measured, the magnitude of the difference in displacement and sd varied depending on speed, leading to inconclusive results in terms of absolute or relative validity of the simulator. again, these results seem to provide strong argument supporting driving simulators as recommended research tools. 2.3. road safety countermeasures the simulator validity for evaluating the relation of a driver with road design and traffic control devices was another issue often examined. riemersma et al. [6] performed two experiments aiming to understand, explain and describe these effects. first, the daimlerbenz driving simulator in which the authors created a scenario where the driver entered a dutch village area. the goal of the simulator was to register changes in desired speed adjustments. the drivers were divided into 2 groups and drove the car either on a normal road or on the same road with special speed reduction objects. three infrastructure changes were implemented in the second scenario, those being a median strip, coloured asphalt and a portal gate. analysis of the vehicle speeds in both scenarios revealed decreased speed values in the scenario with special objects installed. moreover, faster drivers reduced their speeds more than average speed drivers. the second experiment involved both real-world driving and in-simulator ride, utilizing their university driving simulator to perform speed reducing infrastructure testing. the experiment was conducted on 24 participants, of which each drove through the village entrance 12 times. infrastructure changes varied during the rides. driving speed comparison between real-world and simulator revealed higher average speed of in-simulator driving than on road conditions at distance of approx. 400 meters from village entrance. nevertheless, an interesting finding was that, after the 400m distance point, simulator drivers reduced their speed to a greater extend, thus entering the village at a lower speed. the magnitude of speed reduction was evaluated as 3 times bigger for simulator than real-world (25,7% and 8,6% respectively) this can indicate similarly to chapter 2.5.2, relative but not absolute validity of the simulator results. due to these results, the authors concluded that a simulator was an effective tool for evaluating speed reduction methods. following research of mcavoy and datta [16], found in their study that amos ii simulator has contrary results for evaluating night traffic speed control devices in construction zones. a model construction zone on a highway with warning lights was designed and implemented into simulator conditions. 35 drivers took a ride in simulator at the road described above, while in naturalistic study vehicles average speed was measured on road with the use of radar guns. the summary of the results led the researchers to conclude that the simulator failed to provide proof for relative 70 vol. 12/2017 cross-study research of driving simulator or absolute validity standards. a suggesteed explanation for that fact was the subject’s perception of risk associated with driving through real-world work zones at night in a simulator-provided experience. the final paper reviewed in relation to drivers’ behaviour studied using a simulator was the work of godley et al. [13]. in this study again, similarly to riemersma work [6], relative validity of simulator results was found for effect of rumble strips as a speed-reduction method. 2.4. complex driving behaviours another examined validity in short history of driving simulator experiments is in relation to the complex behaviours of drivers such as tasks of divided attention. reed & green [9] investigated validity of the driving simulator constructed at the university of michigan transportation research institute. the goal of the study was to assess decrements in driving performance during a manual car phone task. participants of the experiment were divided into 2 groups: young (2030 years) and old (above 60 years old), and received instructions to dial a manual phone call during vehicle driving tasks both in the on-road setting and in a simulator. conclusions of the work exhibited that speed parameters (sd of speed and throttle position) were similar in both studied conditions, whereas lane positioning (i.e. mean lateral speed and sd of steering wheel position) varied significantly more in the simulator. overall, for seven of the ten variables measured in the experiments, a significant correlation was found between two environments (authors provide correlation factor r in range of 0.43 to 0.76). for all experiment participants regardless of the age group, the phone task resulted in a decrease in driving abilities in terms of speed control and lateral position of vehicle. worth mentioning is the fact that, just like in the previous studies, the magnitude of impact of particular driving conditions was higher in a simulator than on the road. additionally, older drivers seemed to suffer more from the phone dialing task than young drivers in terms of driving abilities. few other studies in this area were performed, but ones worth mentioning are lee, lee and cameron [14] (distinguishing between drivers with different abilities), lee & lee [18] (identifying drivers at risk of future violations), reimer et al. [19] and salvucci et al. [20] who investigated an iphone, button style flip-phone and ipod distraction on driver performance. results of these works however, are based on very few participants therefore, although similar to the ones described above, they will not be summoned in detail here. 2.5. physiological measures the final validity of importance investigated in this research are psychological responses. while the previous sections described a direct way of measuring driving behaviour, physiological responses are commonly used to determine driver awareness and thus produce some useful insight into driver behaviour and reactions across both simulator and real-world contexts. an important condition that is required from driving simulators all over the world is that simulations should induce and provoke similar physiological responses as real-world experience. first research conducted on that topic was watts and quimby [4], who used skin conductance measures for subjects in a transport and road research laboratory simulator. due to the technological limitations of the time, driving simulation was performed in form of film presented to the drivers that was placed to create the impression of being the driver of a vehicle while film from a drive on a road was displayed. the presented film was 30 minutes long, and consisted of 45 hazardous situations for which drivers had to evaluate the perceived danger on a 10-point scale. conductance of skin was measured for all participant during the film. the measurements were then repeated for a real-world driving condition and the results were compared. in the presented results, the risk assessments were comparable for both environments, and moreover, the perceived risk ratings were positively associated with skin conductance levels (r = 0.78). a more recent study performed by slick et al. [21] utilized measures of both skin conductance and heart rate in both simulator and real-world conditions. the experiment performed in drivesafety ds600c motionbased simulator was made of four specific driving tasks performed on both stop sign and intersection traffic lights. the study suggests that there are no significant differences between skin conductance and heart rate measures between the simulator and realworld driving conditions. the only exception for female drivers was that their mean change in heart rate was significantly smaller in the simulated condition than while driving on-road. the most recent study is one analyzed by mullen, charlton, devlin and bedard [3]. bedard in his research measured heart rate and respiration for participants exposed to threatening situations while driving stisim 400 simulator. two intersection scenarios were investigated in depth with the green light turning into orange immediately after driver approached the intersection, and vehicle appearing from the right shoulder of the road in front of the driver. results indicated high measure of presence of the simulator. in each observed situation, the driver’s heart rate increased significantly, and 65% of the participants showed increased respiration rates in risky situations. overall, these studies suggest that physiological measures in general demonstrate the validity of simulators. overall, these studies suggest that physiological measures in general demonstrate the validity of simulators. 3. summary and conclusions many studies have investigated the validity of simulators for examining driving behaviour, leading to a 71 michal matowicki, ondřej přibyl acta polytechnica ctu proceedings conclusion that there is sufficient evidence to suggest that simulators provide a valid tool for assessing various driving performance measures. some important findings on the topic of simulator implementation can be found in studies like [22–24], where overwhelming majority of the studies suggest that there are no obstacles in using driving simulators. these studies suggest that simulators provide approximated behaviour and reactionsof the drivers (relative validity) , despite the fact that it does not replicate on-road behaviour (absolute validity) exactly. the studies under review confirm a moderate to high correlation of the simulator-achieved results for a whole range of cognitive measures of drivers as well as physiological reactions. this implies that a driving simulator is a sufficient tool for majority of research purposes, and thus may be accepted as suitable for the collection of desired data. nevertheless, it is of high importance to remember that the fact that a driving simulator was shown to be valid in a previous setting does not guarantee it will be suitable for the next one. therefore, it is highly recommended to perform additional simulator validation of the results for each specific setting and the scenario designed for a given simulation experiment. this ensures the researchers are capable of utilising this data to create solutions applicable to real-world situations. on road measures will surely be necessary and prevail over the driving simulator method in research situations in which absolute validity of measured variables is required. although there are certain drawbacks and limitations to current simulator deployment in transport research, there is a confident promise of overcoming this issue through technological advancement and, thus, broadening of simulator validity by creating better imitation and perceptual projection of on-road condition for simulator drivers. this will undoubtedly increase the role of simulators as a tool for driving performance measurements as well as for studying driving behaviour. such studies are certainly important not only with respect to the traditional transportation research tasks (transportation planning, study of traffic flow characteristics, optimization of traffic control and others), but also for new and emerging tasks such as within the field of smart cities (resp. smart mobility), study of drivers using cooperative vehicles by přibyl and svítek [25]. the authors of this paper place high importance on understanding which real-life conditions a driving simulator can replicate and aimed to provide guidance in this respect through this document. acknowledgements this article was supported by funds from grant no. sgs16/186/ohk2/2t/16. references [1] e. commission. white paper roadmap to a single european transport area – towards a competitive and resource efficient transport system. tech. rep., european union commission, 2011. date-added: 2011-03-08. [2] f. bella. driving simulator for speed research on two-lane rural road. elsevier accident analysis & prevention 40:1078–1087, 2008. [3] n. mullen, a. charlton, j.and devlin, m. bedard. simulation vlidity: behaviors observed on the simulator and on the road. handbook of driving simulation for engineering, medicine, and psychology. crc press, 2011. [4] g. j. watts, a. r. quimby. design and validation of driving simulator for use in perceptual studies. tech. rep., transport and road research laboratory, 1979. [5] g. j. blaauw. driving experience and task demands in simulator and instrumented car: a validation study. human factors 24:473–486, 1982. [6] a. r. r. riemersma, j. b. j.and van der horst, w. hoekstra, g. m. m. alink, n. otten. the validity of a driving simulator in evaluating speed-reducing measures. traffic engineering and control 31:416–420, 1990. [7] j. tornros. driving behaviour in real world in a real and simulated road tunnel – a validation study. accidents analysis & prevention 30:497–503, 1998. [8] m. g. wade, c. hammonds. simulator validation: evaluating driver performance in simulation and the real world. tech. rep., minnesota department of transportation, 1998. [9] m. p. reed, p. green. comparison of driving performance on-road and in low-cost simulator using a concurrent telephone dialling task. ergonomics 42:1015–1037, 1999. [10] h. klee, c. bauer, e. radwan, h. al-deek. preliminary validation of driving simulator based on forward speed. transportation research record 1689:33–39, 1999. [11] a. c. bittner, o. simsek, w. h. levison, j. l. campbell. on-road versus simulator data in driver model development: driver performance model experience. transportation research record 44:303–313, 2002. [12] e. blana, j. golias. differences between vehicle lateral displacement on the road and in a fixed-base simulator. human factors 44:303–313, 2002. [13] s. t. godley, t. j. triggs, b. n. fildes. driving simulator validation for speed research. accident analysis & prevention 34:589–600, 2002. [14] h. c. lee, a. h. lee, d. cameron. validation of driving simulator by measuring the visual attention skill of older adult drivers. american journal of occupational therapy 57:324–328, 2003. [15] f. bella. validation of driving simulator for work zone design. transportation research record 1937:136–144, 2005. [16] d. s. mcavoy, k. l. schattler, t. k. datta. driving simulator validation for nighttime construction work zone devices. transposrtation research record 2015:55–63, 2007. [17] d. shinar, a. ronen. validation of speed perception and production in a single screen simulator. advances in transportation studies special issue:51–56, 2007. 72 vol. 12/2017 cross-study research of driving simulator [18] h. c. lee, a. h. lee. identifying older drivers at risk of traffic violations by using a driving simulator: a 3-years longitudinal study. american journal of occupational therapy 57:97–100, 2005. [19] b. reimer, b. mehler, b. donmez, et al. a driving simulator study examining phone dialling with iphone vs. button style flip-phone. proceedings of the human factors and ergonomics society annual meeting 56(1):2191–2195, 2012. [20] d. d. salvucci, d. markley, m. zuber, d. p. brumby. ipod distraction: effects of portable music-player use on driver performance. proceedings of the sigchi conference on human factors in computing systems pp. 243–250, 2007. [21] r. slick, d. f. evans, e. kim, j. steele. using simulators to train novice teen drivers: assessing psychological fidelity as a precursor to transfer of training. paper presented at driving simulation conference – tsukuba 2006. [22] p. bouchner, s. novotny, r. pieknik, o. sykora. assessments of grade crossing warning and signalization devices driving simulator study. pp. 179–183. 2011. cited by 0. [23] p. bouchner, s. novotny. development of advanced driving simulator: steering wheel and brake pedal feedback. pp. 170–174. 2011. cited by 0. [24] a. mashko, p. bouchner, d. rozhdestvenskiy, s. novotny. virtual traffic signs – assessment of an alternative adas user interface with use of driving simulator. advances in transportation studies 1:37–50, 2016. cited by 0. [25] o. pribyl, m. svitek. system-oriented approach to smart cities. proceedings of the first ieee international smart cities conference isbn 978-1-4673-6551-2 2015. 73 acta polytechnica ctu proceedings 12:68–73, 2017 1 introduction 2 reviewed studies 2.1 speed adjustment 2.2 lateral position of vehicle 2.3 road safety countermeasures 2.4 complex driving behaviours 2.5 physiological measures 3 summary and conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0039 acta polytechnica ctu proceedings 11:39–42, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app its on railway martin leso czech technical university in prague, faculty of transportation sciences, praha, czech republic correspondence: leso@fd.cvut.cz abstract. development of technology management and security, transmission and information systems allows the concept of process management and security of railways and organizing the operation of a new, as yet non-phase in a non-traditional way. requirements for higher efficiency, accuracy and flexibility of the attractiveness of rail transport connection in the integrated transport systems are a challenge for the railways. these requirements can be solved just by changing the concept and technology of the presented here as an its-r. keywords: ertms/etcs, gsm-r, inteligent traffic systems on railway (its-r), traffic management systems (tms), railway control and command systems (ccs). 1. state of the art modern technologies are applied for control, command and signaling and in the area of information systems (for both control and passengers check-in) on the czech railway. electronic systems are installed (station electronic interlocking systems, crossing interlocking system and etc.) including remote control systems (central dispatching center). information systems allow to enter process and provide data to be able to control direct and on strategic level control and plan of railway transport. systems also provide information to passengers via panels and monitors. today’s approach is focused towards to install interlocking technologies on the infrastructure and the problems related to the movement of vehicles are missed out. vehicles are equipped by systems checking the driver’s watchfullness and by the lvz system transmiting signals to the driver’s cabine. but, the responsibility of reading signals and satisfy all rule remains on the driver himself. about 1500 km of the czech corridors is equipped by the system lvz. remaining parts of the infrastructure have no lzv and the driver is monitored only by periodical vigilance button pushing. this state of the art is analogical in easterneuropean countries (poland, slovakia, hungary). some vehicles (čd 471 series) are equipped by intelligent technology ato (automatic train operation), denoted avv in czech language, that makes possible to automate vehicle and driver’s work control. these systems lead to decrease operating costs but their wide application is not optimal now due to missing pieces of information from the infrastructure and absence of safety level of vehicle driving. despite the avv system developed in the czech republic reached excellent operating parameters on the world level. ertms implementation plan was ratified and published in 2014. this plan is a crucial document to define the ertms implementation strategy in the czech republic. the fundamental change (in comparison with existing plans) is a new strategy to equip vehicles by "etcs onboard part". on the continuous track sections equipped with etcs of hundreds of km length immediately after entry into service will be prohibited trains without functional compatible vehicle on-board etcs. it is planned to equip about 1000 vehicles by etcs onboard system up to 2020 and about 1500 vehicles up to 2024. it may be stated the most vehicles will be equipped by etcs. but ertms implementation plan is concerned only to ten-t track net, other track net is not supported yet. the czech rail net included core ten-t track with length of 1329 km (3670 km of rails), others tracks ten-t 1265 km (3022 km), other main tracks of the length 2430 km (3633 km kolejí) and 4409 km (5079 km kolejí) of regional tracks. european legislative tsi (technical specifications of interoperability) defines requests to adapt main and regional rails outside of ten-t including all vehicles and industrial vehicles to be interoperable, i.e. etcs compatible. the issue of dealing with these types of lines forming around 2 / 3 of railway network in czech republic, therefore remains a major challenge for the future. the ertms implementation plan prepare the implementation of its-r systems well. with regard to the locomotives complete on board etcs fulfilment, both national main line and regional lines shall be calculated to be vehicles to be equipped with ertms / etcs on-board units. the assumption of full ertms/etcs onboard fulfillment is the big opportunity for change scope and type of interlocking technologies installed on infrastructure. today’s "full" installation with railway interlocking systems (station, track and level crossing systems with signals, track circuits od axel counters and etc.) on this types of tracks aren’t full installed not yet. on focus mainly to the regional tracks, using of conventional interlocking systems is not economical justified. these types of track are therefore suitable for installation of ertms/etcs systems, mainly 39 http://dx.doi.org/10.14311/app.2017.11.0039 http://ojs.cvut.cz/ojs/index.php/app martin leso acta polytechnica ctu proceedings etcs – level 1, or etcs – ls mode. most forward-looking of all functions with regard to its systems may be etcs – l3 (currently validated as ertms / etcs regional). with these systems may also consider the use of new technologies gnss, which are not yet in place in the tsi, but there may be early implemented. on this types of tracks, on the corridors main tracks to, it is possible to minimalize investment and services costs by choosing the strategy and type of control of the railway system based on minimalizing of parts of infrastructure technologies with reduction of signals and vehicle detection systems. it is necessary to consider, having regard to its applicability of technologies, the use of radio communication between vehicle and infrastructure. they must also be addressed, in particular freight train integrity to enable the detection systems in infrastructure. in 2015 was also approved and published other strategic document of ministry of transportation of czech republic (md čr), the "action plan its", which also deals with the implementation of its to rail. action plan for ertms / etcs and its implementation in the czech republic and, in principle, only referred to the downstream transport technologies (remote control, automation driving, etc..). the overall system is the concept of its in rail. in another development, it is therefore advisable to ertms / etcs, gsm – r and its one unit, in order to implement appropriate design and integration of all systems to ensure the function of its – r. 2. european best practices western european countries, the issue of long-term facilities of vehicles already addressed since 1950 mainly due to the provision of safe movement of vehicles. for example, german railway has wholly used vehicle systems controlling train movements (lzb main lines, branch lines pzb). very high level of amenities vehicles ertms / etcs (level ls, level 2) currently have swiss railways. for example, german railways has a wholly owned equipment vehicles systems controlling train movements (main line system lzb, pzb byline). very high level equipment of vehicles ertms / etcs (level ls, step 2) currently has swiss railways. in europe and globally range, can be seen guideline facilities especially vehicle management systems for rail traffic on the principles of ertms / etcs. ertms / etcs is promoted by the european union as one of the subsystems to ensure the interoperability of the whole rail system in europe. ertms / etcs ensures compliance with the requirements of the its-r in particular parameters ensure the safety of vehicle movement. other parameters, such as reliability and operational efficiency, are more dependent on the conditions of application of the system in specific countries. a comprehensive approach to its technology is not yet defined, but it is limited to defining the figure 1. interlinkages partial rail subsystems in its-r. structure of the exchange of data on passenger and freight (taf, tap) – subsystem of telematics applications for freight and passenger transport. in the area of intelligent transport systems can be seen similar trends as in the czech republic, ie. centralizing operations management together with the strong support of information systems for traffic management, but at the european level is already increasingly applied in real operation automatic setting routes, including optimizing traffic management (ie. the swiss system rcs). apply new systems for the management and security of railway traffic. 3. system approach to railways the subject of deployment of the railway is an intelligent transport system whose architecture and elements are interconnected in order to optimize and increase the efficiency of the whole process (fig. 1). its overall concept must be focused on increasing the parameters of the rail system, particularly the safety, reliability, economic efficiency operation. assumption for the goals of its on the rail is availability of all vehicles as a main target in the system, which provides transport (goods and people). these include the installation of ertms / etcs system including automatic train control, wireless data transfer by train (gsm-r). these basic components enable implementation of the concept of intelligent vehicle concept, which ensures the function as safe motion control, as well as optimizing its movement. for proper operation of intelligent vehicles, it is necessary to equip the corresponding intelligent infrastructure that will provide adequate information, intelligent vehicles with the least economic and technical complexity. a necessary condition is also coverage of wireless twoway communication infrastructure, which can provide information not only for the management and operation, but it will also provide an information service for dispatching and public services to passengers. target status should lead to the integration and overall optimal solutions to the infrastructure, and 40 vol. 11/2017 its on railway figure 2. conceptual vision of its-r. especially on vehicles with the aim of using all the components for securing the services of its-r. to ensure sufficient functionality and quality of services must be technology-r its implemented using modern telecommunications and emerging technologies. as a basis for solving the safety features of the system (of trains) must be the movement of all vehicles within the control signaling system ertms / etcs (fig. 2. part draw black). these systems need infrastructure equipment gsm-r system. according to the chosen level of ertms / etcs infrastructure must be equipped with the appropriate infrastructure components. for etcs-l1 (ls) and etcs-l2 infrastructure must be fully equipped with conventional technology infrastructure. etcs level-l3 enables significant savings on the side of infrastructure facilities, conventional systems management and security are not necessary and are fully replaced by technology on the vehicle condition to ensure the integrity (integrity) of the train. to increase efficiency and optimize operations management, increase infrastructure capacity, reduce operating costs (fuel consumption and energy use, vehicle wear and infrastructure) is required all vehicles equipped with automatic train control system. this subsystem most of the information for their activities (information on the line, the vehicle allowed to travel etc ..) can be obtained through a signaling system ertms / etcs. to ensure greater attractiveness of traveling using modern information and check-in systems and logistics services for parent carriers and their customers is also needed to ensure coverage of railway infrastructure, high-speed data services (eg. the technology 4g / 5g; fig. 2. showing green and red). high speed data connectivity is to be ensured in all vehicles, i.e. not only driving the vehicle but also to all connection (in the case of complete sets embedded) vehicles. 4. advantages of the benefits of its deployment at railway the introduction of intelligent transport systems on the railways should provide ways to increase the efficiency of the entire transport system in the czech republic: (1.) state managers of transport infrastructures: (a) increasing the capacity of transport infrastructure (b) increasing transport reliability, minimize delays (c) reducing wear of the transport infrastructure (d) efficient rail transport with possibility of creation the integration transport systems with using railways as the transport backbone (e) optimization of investment in transport infrastructure (f) increasing the level of safety throughout the rail network (thanks to the strategy is to increase the safety of the entire network, after which the vehicle is moving under conditions appropriate infrastructure facilities) (2.) railway carrier and passangers: (a) a. reduction of vehicle operation costs (energy, operating costs) (b) reducing of traffic accidents (c) increasing the attractiveness of rail transport for passengers (d) increasing the attractiveness of rail transport for freight transporters (e) increasing the attractiveness of rail transport for carriers (passenger and freight) (f) increased profits from transport (3.) railway industry: (a) an opportunity for the development and implementation of new products and services (b) exports of modern technology with the potential to assert themselves on foreign markets 5. final conclusion this paper describes the concept of new technology for the management and security of railway traffic solutions enable intelligent transportation systems. at present day, all the key technologies are developed yet. if we want to fulfill the above-mentioned strategic action is needed as soon as possible to start with a pilot project of the its-r. the pilot project should be implemented at the appropriate regional railway track, which will be verified and mastering functionality of the various sub-components and the whole concept of traffic management. the subject of verification should be technology ertms / etcs l3 system including automatic train control, information and communication systems to ensure transmission of data necessary for the management and operation, logistic processes and services for passengers. after this necessary phase of technology adoption may be 41 martin leso acta polytechnica ctu proceedings considered for their further expansion into the real network traffic lanes regional and national railways. the main benefit of technology its-r is not only to increase safety on the regional railways, but also increase the flexibility and the possibility of significant automation and optimization of traffic on main line of railways. consequences should be in greater competitiveness and attractiveness of rail transport. 42 acta polytechnica ctu proceedings 11:39–42, 2017 1 state of the art 2 european best practices 3 system approach to railways 4 advantages of the benefits of its deployment at railway 5 final conclusion acta polytechnica ctu proceedings doi:10.14311/app.2017.9.0001 acta polytechnica ctu proceedings 9:1–7, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of parameters of surface integrity of machined duplex and austenite stainless steels in relation to tool geometry jiří čapeka, ∗, zdeněk pitrmucb, kamil kolaříka, libor beránekb, nikolaj ganeva a department of solid state engineering, faculty of nuclear sciences and physical engineering, czech technical university in prague b department of machining, process planning and metrology, faculty of mechanical engineering, czech technical university in prague ∗ corresponding author: capekjir@fjfi.cvut.cz abstract. the goal of this contribution was to describe parameters of surface integrity of two machined materials; austenite and duplex stainless steel. residual stresses and presence of straininduced martensite were studied as a function of the side rake angle. residual stresses of surface and sub-surface layers were determined using x-ray diffraction techniques and hole-drilling method. by using x-ray diffraction, it is possible to determine residual stresses in each phase separately, in comparison with hole-drilling method. the presence of strain-induced martensite was investigated using barkhausen noise and optical microscope. keywords: stainless steels, machining, x-ray diffraction, residual stresses, strain-induced martensite. 1. introduction actual state of technology of sewage disposal plants (sdp) basically does not solve the problem of purification of fats and oils. using mobile tricanter device substantially changes the attitude of waste processing to sdp. it removes transporting of waste from small, problematic, or dangerous environmental sources and enables using device on the polluted place. the mobile version of tricanter device poses increased requirement on reducing of energy consumption and mobility given by size and weight reducing. the mobile separation tricanter may not be discernible from mobile separation decanter device. the differences are mainly in centrifuge and accessories of device. tricanter device enables three-phase separation with simultaneous separation of two immiscible liquids with different densities and one solid phase, provided that the solid phase is the heaviest phase using high revolution, i.e. centrifugal force. the main problems are removing of individual phases and purity observation. duplex and austenite steels are usually used for their properties, primarily for the high corrosion resistance. duplex steels combine properties of both phases and moreover, due to two-phase microstructure, some properties are better than high-alloyed austenite steel, e.g. abrasion resistance [1]. abrasion and corrosion resistance have to be one of the main required properties. not only properties of used material but it is necessary to investigate quality of machined surface — surface integrity, too [2, 3]. both, austenite and duplex steels have relatively low thermal conductivity (approx. 16 w/mk) [4] which leads to insufficient heat distribution into chip and workpiece and to excessive heat accumulation in cutting zone. this heat generation can result in microstructural changes, local changes of chemical composition, surface discoloration or inducing undesirable tensile residual stresses. the gradient of residual stresses (rs) is more important indicator of machined material than surface rs. a situation may arise that there are the favorable compressive rs on the surface but with a steep rs gradient which can result in very high unfavorable tensile rs in the sub-surface layers, and vice versa [5]. for this reason, it is very important to investigate the rs gradient. analysis of polycrystalline materials by x-ray diffraction methods is suitable for gaining information about state of rs of both the surface and subsurface layers. on the other hand, the other methods as hole-drilling determine the rs gradient from total relieved deformation of material after disruption of rs balance. the strain-induced martensite can be caused by sufficiently high rate of plastic deformation of surface layers and also by temperature instability of austenite [6, 7]. martensite is harder and more brittle in comparison with austenite steels and thus changes their properties. for strain-induced martensite determination, e.g. the optical microscopy or barkhausen noise analysis can be used. 2. theory duplex stainless steels have high corrosion resistance in many environments, where the standard austenite steel is consumed, and where its properties sig1 http://dx.doi.org/10.14311/app.2017.9.0001 http://ojs.cvut.cz/ojs/index.php/app j. čapek, z. pitrmuc, k. kolařík et al. acta polytechnica ctu proceedings nificantly exceed austenite steel. thereby, smaller amount of material from duplex steel is necessary to manufacture function components. austenite and duplex steels are susceptible to mechanical reinforcement, i.e. local changes in mechanical properties of surface layers. local changes, e.g. hardness, can lead to tools vibration during machining of the final component, which results in additional material inhomogeneity and blunting tool [8]. realizing that austenite steel has face centred cubic (fcc) lattice with close-packing structure of atoms, the primary slip system goes along {111}〈110〉 the number of slip systems is 12, which is the sufficient amount to plastic deformation. moving dislocations form so called stair-rod dislocations which have small stacking fault energy, i.e. high energy is necessary to have for intersect or cross slip of these dislocations [9]. therefore, the austenite steels are prone to workhardening, which cause mechanical modification and inhomogeneity on the machined surface, and leads to e.g. unstable chip formation. the second type of deformation of high alloyed austenite steel takes place via twins. on the contrary, the ferrite crystallizes in a body centered cubic lattice (bcc). the direction slip in bcc materials is always 〈111〉. since in the bcc lattice is not close-packing structure of atoms, more slip planes assert during the deformation, mostly planes {110} and {211}. in engineering practice, the residual stresses are often determined on the basis of total sample deformation (e.g. hole-drilling method). this procedure considers the solid state as a compact continuous body, and therefore could not take into account degradation processes which rise separately into each phase components. using x-ray diffraction (xrd) method, the residual stresses can be determined into both the phases separately [10]. for strain-induced martensite determination, several method can be used. x-ray diffraction and barkhausen noise belong to non-destructive techniques and e.g. microhardness measurements or optical microscopy to destructive analysis. 3. experiment the tested samples of tube shape of 100/86 mm in diameter were made of aisi 304 (austenite) and aisi 2205 (duplex) type of stainless steel. the samples were annealed in air laboratory furnace for 5 hours at 420 ◦c in order to reduce bulk macroscopic residual stresses. for machining of the surfaces, four types of side rake angle were used (−6◦; −2◦; +7◦ and +12◦). side rake angles are considered in combination with particular insert holder, which has negative rake angle (−6◦). dclnr/l r-clamp tool-holders with lead angle of 95◦ (side cutting edge angle −5◦) for four 80◦ negative rhombic inserts were used, namely f3m, sf, nf, and pp chip breakers of iscar cutting tools. all inserts had the same tip radius 0.4 mm. for elimination of blunting tool effect, the cutting tool was always new for machining of each tube segment. the surface of the austenite sample was machined with different tool’s feed, f = 0.1, 0.15; 0.2 and 0.31 mm/rev, using constant value of cutting depth 0.5 mm and cutting speed 155 m/min. the second series of machined samples were set up the cutting conditions of experiment using different tool geometry: feed rate 0.14 mm/rev, cutting speed 140 m/min, and depth of cut 2 mm. direction of feed rate was parallel to axis of the sample (tube) a and perpendicular to tangential direction t. according to the principles of design of experiments (doe) method, three 1cm tube segments were machined using the same cutting conditions. using mnkα and crkα radiation, x’pert pro mpd diffractometer was used to measure lattice deformations in austenite and ferrite, respectively. the average penetration depth of x-ray radiation is approx. 4 µm and 6 µm for ferrite and austenite phase, respectively. diffraction angles 2θhkl were determined from the peaks of the diffraction lines kα1 of planes {311} and {211} of austenite and ferrite, respectively. diffraction lines kα1 were fitted by pearson vii function and rachinger’s method was used for separation of the diffraction lines kα1 and kα2. for residual stress determination, winholtz & cohen method [11] and x-ray elastic constants 1/2s2 = 7.18 tpa−1, s1 = −1.20 tpa−1 and 1/2s2 = 5.75 tpa−1, s1 = −1.25 tpa−1 were used for austenite and ferrite phase, respectively. in order to analyze the stress gradients beneath the samples surface, layers of material were gradually removed by electro-chemical polishing in the center of the sample. hole-drilling method was performed using sintered carbide milling drill of 1.8 mm in diameter and the holes had depth of 2 mm. the detection of released deformations was done by 3 rectangular tensometric rosettes. the stresses were calculating using macroscopic elastic constants: young modulus 200 gpa and poisson ratio 0.3. the measurements of barkhausen noise were performed using a commercial unit stresstech rollscan 350 magneto-elastic analyser with a standard sensor. main parameters of the applied method were: sinusoidal shape of magnetic signal, magnetic voltage 3.5 v, and frequency 75 hz. in practice typical expectable penetration depth in this experimental arrangement is in the range of 20 µm, but generally depend on frequency and the analysed material. for obtaining optical microscopy images, the radial cut of each sample was made. the electrolytic etching with 10% oxalic acid of investigated surfaces was employed. the images were achieved by inclined illumination with 250× magnification. 2 vol. 9/2017 parameters of surface integrity of duplex and austenite steels figure 1. averaged values of residual stresses 〈σa〉, mpa (top) and 〈σt〉, mpa (bottom) of both phases depending on tool’s feed f, mm/rev. 4. results and discussions 4.1. dependence on tool’s feed f in figs. 1, there are dependence of residual stresses values of ferritic and austenitic phases on the tool’s feed. the dependence of residual stresses 〈σ〉 on the tool’s feed f is increasing, see figs. 1. this trend is identical in the both analyzed directions; axial and tangential. it can be concluded that using a smaller tool’s feed leads to lower values of residual stresses. this effect is related to "an amount of chip". higher tool’s feed, cutting depth and cutting speed lead to higher cutting force and higher heat generation. if the heat effect is dominant, the tensile stresses are generated. generally, the temperature influence causes the tensile rs and contrarily, the plastic deformation leads to compressive rs. the type of the rs and their value deeply depend on the mechanical and thermal properties of the machined material [5, 12]. 4.2. dependence on the side rake angle 4.2.1. surface residual stresses in figs. 2, there are influences of surface macroscopic residual stresses 〈σa〉, 〈σt〉, mpa on the side rake angle, ◦. these residual stresses were averaged from three values of rs of tube segments machined using same side rake angle. generally, the increasing of the side rake angle in the positive direction leads to a lowering of cutting force and temperature in the cutting zone [5]. for prediction of rs dependence on the side rake angle, the yield strength ratio rm/rp0.2 of the given material is necessary to take into account. for austenite steel, the yield strength ratio is approx. 2.5, which is typical value for plastic material. the tensile rs are created during machining using great load on the cutting tool, i.e. using the negative side rake angle, and smaller load causes the compressive rs, i.e. using the positive side rake angle [5]. for this reason, higher compressive (axial direction) and smaller tensile (tangential direction) rs were determined with increasing of the side rake angle, see fig. 2a. on the other hand, for ferrite steel, the yield strength ratio is less than 1.25, which is typical value for elastic materials. according to [5], the shear type chips should be created which should cause the interruption of the connection between chips and the material. the additional effect of strain filed of chips is not transferred in the machined surface. for this reason, the greater force causes that the plastic deformation influence is predominant and higher compressive or smaller tensile rs may be determined with increasing of the side rake angle. furthermore, for duplex steel, which consists of both phases, it is possible to presume that the dependence of rs on the side rake angle is generally not monotonic for both the phases because of their mutual influence during plastic deformation, see figs. 2b, 2c. 4.2.2. residual stresses gradient the rs gradients depending on the side rake angles were determined, see figs. 3. as can be seen in figs. 3a, 3b, the rs gradients are without any dependence on the side rake angle in the case of austenite steel. on the contrary, for duplex steel, there are clearly differences of the rs values depending on side rake angle in both austenite and ferrite phase. the reason resides in two phase material and mutual influence of both phases during plastic deformation. in the case of one phase steel, the difference of plastic deformation power is not so significant to change the state of rs in the sub-surface layers using the different side rake angle, see figs. 3a, 3b. on the contrary, for two phase steel, the differences of rs are evident, see figs. 3c, 3d. with the increasing side rake angles, the maximum stress position is moved deeper into the material which is in line with results in [13]. in figs. 4, there is a comparison of rs gradients determined by xrd and hole-drilling method. mostly, the macroscopic rs of bulk material can be estimated as approx. zero, see austenite steel in figs 3a, 3b. nevertheless, the compressive rs of ferrite and the tensile rs of austenite phase of bulk are seen in figs. 3c, 3d. by xrd, the determination of the rs is possible in each phase, separately. because of the similar mass ratio of ferrite and austenite phase in the duplex steel, the rs of bulk can be approximately predicted by value σbulk ≈ (σf errite + σaustenite)/2. however, 3 j. čapek, z. pitrmuc, k. kolařík et al. acta polytechnica ctu proceedings (a) . austenite steel. (b) . duplex steel — austenite phase. (c) . duplex steel — ferrite phase. figure 2. axial and tangential residual stresses 〈σa〉, 〈σt〉, mpa on the side rake angle, ◦. the hole-drilling method is based on deformation of the measured material as the whole. it is evident from fig. 4 that σbulk ≈ σhole−drill ≈ 0 which is in line with expectation [10]. moreover, the continuity of both methods of rs determination was verified. 4.2.3. strain-induced martensite determination in fig. 5, there are the optical microscopy images of the radial cut for two types of chip breakers, for maximum negative (f3m chip breaker, i.e. side rake angle −6◦ ) and maximally positive (sf chip breaker, i.e. side rake angle +12◦ ). from theory follows that the primary shear zone is decreasing with increasing positive side rake angle. in other words, increasing side rake angle results in lower cutting force and temperature [14]. due to (a) . austenite steel — axial direction. (b) . austenite steel — tangential direction. (c) . duplex steel — austenite phase. (d) . duplex steel — ferrite phase. figure 3. residual stress gradients σa,σt using side rake angles −6◦, −2◦ and +12◦. 4 vol. 9/2017 parameters of surface integrity of duplex and austenite steels (a) . axial direction; side rake angle −6◦. (b) . tangential direction; side rake angle −6◦. (c) . axial direction; side rake angle +12◦. (d) . tangential direction; side rake angle +12◦. figure 4. residual stress gradients σa,σt determined by xrd and hole-drilling method. lower cutting force, the primary shear zone tends to get wider and the plastically affected surface layers is thinner. in figs. 5a and 5b, there is evident the possibility presence of strain-induced martensite. for confirmation of this statement, other analysis is necessary to do, e.g. barkhausen noise analysis, see fig. 6. after confirmation of presence of strain-induced martensite, it is evident that there are differences between thickness of surface layer with strain-induced martensite, see figs. 5a and 5b. increasing positive side rake angle, the thickness of mentioned layer is decreasing which is in line with presented theory. in the case of duplex steel, see figs. 5c and 5d, it is possible to observe differences in amount of straininduced martensite, i.e. dark blue grains. with increasing positive side rake angle, the amount of straininduced martensite decreases. using higher positive side rake angle, the cutting tool causes lower plastic deformation and therefore, the lower amount of straininduced martensite is created. the thickness of layer containing strain-induced martensite is approximately identical. in fig. 6, there is a comparison of barkhausen noise records of machined surface with and without straininduced martensite of austenite steel. the presence of ferromagnetic phase in paramagnetic austenite, i. e. strain-induced martensite, was absolutely confirmed by barkhausen noise. the barkhausen noise record of non-machined surface has constant value, i. e. equal to signal creating by measuring probe. the confirmation of the presence of strain-induced martensite in the duplex steel is more complicated due to presence ferromagnetic ferritic phase. the higher value of magneto-elastic parameters are not only due to higher ratio of ferromagnetic phase but also due to residual stresses, dislocation density, grain-boundaries etc. 5. conclusions the present study showed: • regarding dynamical loading of the final product, it is necessary to have a suitable state of residual stresses in the surface layers, i.e. high value of compressive stresses. for this reason, lower tool’s feed is more preferable when the final product is manufactured. with respect to our results, see figs. 1, and recommendation of tools producer, the value of tool’s feed 0.14 mm/rev was selected. • the surface rs distribution is dependent not only on the side rake angle but on the material, too. for austenite (one phase steel), the dependence of rs on the side rake angle is decreasing. on the contrary, for austenite and ferrite (in the two phase steel), the dependence is not monotonous. the reason is yield strength ratio which is different for austenite and ferrite. • the rs gradients are without any dependence on the side rake angle in the case of austenite steel. 5 j. čapek, z. pitrmuc, k. kolařík et al. acta polytechnica ctu proceedings (a) . austenite steel machined by f3m chip breaker, i.e. side rake angle −6◦ . (b) . austenite steel machined by sf chip breaker, i.e. side rake angle +12◦ . (c) . duplex steel machined by f3m chip breaker, i.e. side rake angle −6◦ . (d) . duplex steel machined by sf chip breaker, i.e. side rake angle +12◦ . figure 5. optical microscopy images of the radial cut of four samples; magnification 250×, electrolytic etching 10% oxalic acid. figure 6. barkhausen noise records of austenite steel of machined surface with and non-machined surface without strain-induced martensite. on the contrary, for duplex steel, there are clearly differences in rs values depending on side rake angle in both austenite and ferrite phase. the reason is mutual influence of the both phases during plastic deformation. • for the similar mass ratio of ferrite and austenite in the duplex steel, the rs determined by hole-drilling method are approx. the average between the rs determined by xrd of ferrite and austenite phase, separately. • moreover, the rs of austenite phase are generally more tensile in comparison with the rs of ferrite phase; see figs. 4, which is in correlation with [5]. • in figs. 5a and 5b, there is not evident difference between twins and strain-induced martensite. for confirmation of strain-induced martensite presence in the austenite steel, the barkhausen noise analysis was used, see fig. 6. • it follows from figs. 5 that with increasing positive side rake angle, the amount of strain-induced martensite and the layer containing strain-induced martensite is decreasing for duplex and austenite steel, respectively. • due to correlation between amount of straininduced martensite and microhardness, it is possible to presume the dependence of microhardness on side rake angle. acknowledgements j. čapek acknowledges support of the grant agency of the czech technical university in prague, grant no. sgs16/245/ohk4/3t/14. others were supported by ta04020658, the technology agency of the czech republic. references [1] r. dakhlaoui, et al. mechanical properties of phases in austeno-ferritic duplex stainless steel—surface stresses studied by x-ray diffraction. mater sci eng: a 444(1):6–17, 2007. doi:http://dx.doi.org/10.1016/j.msea.2006.06.074. 6 http://dx.doi.org/http://dx.doi.org/10.1016/j.msea.2006.06.074 vol. 9/2017 parameters of surface integrity of duplex and austenite steels [2] d. y. jang, et al. surface residual stresses in machined austenitic stainless steel. wear 194(1):168–173, 1996. doi:10.1016/0043-1648(95)06838-4. [3] r. mšsaoubi, et al. residual stress analysis in orthogonal machining of standard and resulfurized aisi 316l steels. j mater process technol 96(1):225–233, 1999. doi:http://dx.doi.org/10.1016/s0924-0136(99)00359-3. [4] a. boniardi, et al. stainless steels, 2011. [2016-06-09], http://www.lucefin.com/wp-content/files_mf/ stainlesssteels_low.pdf. [5] f. neckář, i. kvasnička. vybrané statě z úběru materiálu. české vysoké učení technické, 1991. [6] p. haušild, k. kolařík, m. karlík. characterization of strain-induced martensitic transformation in a301 stainless steel by barkhausen noise measurement. mater des 44:548–554, 2013. doi:http://dx.doi.org/10.1016/j.matdes.2012.08.058. [7] h. t. chen, m. f. yan, s. s. fu. martensite transformation induced by plasma nitrocarburizing on aisi304 austenitic stainless steel. vacuum 105:33–38, 2014. doi:http://dx.doi.org/10.1016/j.vacuum.2014.01.032. [8] j. čapek, et al. residual stress analysis into each phase of turning duplex steel and their influence on service life of rotary and dynamically loaded components. in české vysoké učení technické (ed.), the 5th student scientific conference on solid state physics, pp. 11–15. 2005. [9] j. j. moverare, m. odén. deformation behaviour of a prestrained duplex stainless steel. mater sci eng: a 337(1):25–38, 2002. doi:http://dx.doi.org/10.1016/s0921-5093(02)00022-9. [10] k. kolařík, et al. combining xrd with hole-drilling method in residual stress gradient analysis of laser hardened c45 steel. adv mater res 996:277–282, 2014. doi:10.4028/www.scientific.net/amr.996.277. [11] u. welzel, et al. stress analysis of polycrystalline thin films and surface regions by x-ray diffraction. j appl crystallogr 38(1):1–29, 2005. doi:10.1107/s0021889804029516. [12] t. leppert, r. l. peng. residual stresses in surface layer after dry and mql turning of aisi 316l steel. produc eng 6(4–5):367–374, 2012. doi:10.1007/s11740-012-0389-3. [13] p. dahlman, et al. the influence of rake angle, cutting feed and cutting depth on residual stresses in hard turning. j mater process technol 147(2):181–184, 2004. [14] h. saglam, f. unsacar, s. yaldiz. investigation of the effect of rake angle and approaching angle on main cutting force and tool tip temperature. int j mach tool manu 46(2):132–141, 2006. doi:http://dx.doi.org/10.1016/j.ijmachtools.2005.05.002. 7 http://dx.doi.org/10.1016/0043-1648(95)06838-4 http://dx.doi.org/http://dx.doi.org/10.1016/s0924-0136(99)00359-3 http://www.lucefin.com/wp-content/files_mf/stainlesssteels_low.pdf http://www.lucefin.com/wp-content/files_mf/stainlesssteels_low.pdf http://dx.doi.org/http://dx.doi.org/10.1016/j.matdes.2012.08.058 http://dx.doi.org/http://dx.doi.org/10.1016/j.vacuum.2014.01.032 http://dx.doi.org/http://dx.doi.org/10.1016/s0921-5093(02)00022-9 http://dx.doi.org/10.4028/www.scientific.net/amr.996.277 http://dx.doi.org/10.1107/s0021889804029516 http://dx.doi.org/10.1007/s11740-012-0389-3 http://dx.doi.org/http://dx.doi.org/10.1016/j.ijmachtools.2005.05.002 acta polytechnica ctu proceedings 9:1–7, 2017 1 introduction 2 theory 3 experiment 4 results and discussions 4.1 dependence on tool's feed f 4.2 dependence on the side rake angle 4.2.1 surface residual stresses 4.2.2 residual stresses gradient 4.2.3 strain-induced martensite determination 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.10.0034 acta polytechnica ctu proceedings 10:34–37, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app fiber reinforced sand backfill for underground pedestrian crossing andrey b. ponomarev∗, v.i. kleveko, o.v. moiseeva, k.r. kashapova perm national research polytechnic university, kuibysheva street 109, perm, russia ∗ corresponding author: spstf@pstu.ru abstract. crosswalk is an essential element of the transport system of any city. a priority assignment in any such crossing is to ensure the safety of pedestrians. underground pedestrian crossings are much safer than crossings at another level, but they lose in attraction when it comes to cost. the cost of the construction of an underground pedestrian crossing often leads to project abandonment. the cost of construction can be reduced through the use of flexible corrugated metal structures instead of ordinary concrete structures. the stress-strain state of the structural shells need to be known to facilitate their rational design. the stress-strain state of flexible corrugated metal shells largely depends on the strength and deformation characteristics of the surrounding soil. therefore, improving the characteristics of backfill soil is an urgent task in reducing the cost of construction for such tunnels. one way to improve the strength and deformation characteristics of soils is the use of reinforcement. currently, there are a large number of reinforcement schemes and also associated reinforcing materials. one of the most prospective methods of soil reinforcement is the use of fibre filaments. fibre reinforced soils have significantly higher strength and deformation characteristics in comparison with unreinforced soils. numerical modelling of a tunnel shell made out of a corrugated metal structure was undertaken to evaluate the effectiveness of using fibre reinforced sand. ordinary sand and sand reinforced with polypropylene fibres have been used as soil backfill. the calculation results for a pedestrian tunnel structure involving different strength and deformation characteristics of the backfill soil are presented in this article. keywords: fibre reinforced soil, fem, tunnel lining. 1. introduction crosswalk is an integral element of the transportation system of any city. the first and foremost function of a pedestrian crossing is user safety. annually, in russia, a large number of people travelling on foot are killed, which indicates some lack of effectiveness in the transfer and transport system as a whole [1, 2]. the most commonly used and the most projected type is a pedestrian crossing with a light regulation of type “zebra”. this type of transit actually maximises the probability of pedestrian-car contact. underground pedestrian crossings offer a significantly safer transit to light regulation, but lose on the cost criterion. this cost of construction often leads to abandonment of the concept. however, the cost of construction can be reduced through the use of corrugated metal structures instead of the standard concrete shell [3–6]. to reduce the earthworks and consequently the cost of construction, the pedestrian tunnel is laid at the ground-level for traffic vehicles and in this case is carried out using the device of the mound (fig. 1). earlier studies showed that the optimal cross section is a round pipe [6]. the stress-strain state of flexible metal corrugated shells largely depends on the strength and deformation characteristics of the surrounding soil [7, 8]. therefore, the characteristics of the soil backfill gives an actual opportunity to reduce the cost of construction of pedestrian tunnels. one way to improve the strength and deformation characteristics of soils is the use of reinforcement. currently, there are a large number of schemes for reinforcement and their related reinforcing materials. 2. fibre reinforced sandy backfill one of the most promising ways of reinforcement of soils is using fibre. fibre reinforced soils have a significantly higher strength characteristics compared to unreinforced soils [9–12]. fig. 1 shows the structure of the pedestrian crossing at ground level. this design of pedestrian crossing does not need light regulation, and thus saves time for drivers and pedestrians, requires less effort compared to aboveground and underground pedestrian crossings and generally serves to enhance the safety of all road users. 34 http://dx.doi.org/10.14311/app.2017.10.0034 http://ojs.cvut.cz/ojs/index.php/app vol. 10/2017 fiber reinforced sand backfill figure 1. the construction of a pedestrian crossing at ground level figure 2. a round pipe figure 3. specifications of corrugated sheet gl 55 to identify the optimal soil backfill there was calculation by the finite element method, implemented in the software package plaxis 2d. the construction of a pedestrian crossing at ground level was made for specific geological conditions. at the construction site there were identified the following primers: • loam, layer thickness h = 7.2 m, intercept cohesion c = 20.5 kpa, angle of internal friction ϕ = 18°, modulus of deformation e = 11 mpa; • sand, h = 1.0 m, c = 2 kpa, ϕ = 18°, e = 11 mpa; • gravel, h = 1.8 m, c = 1 kpa, ϕ = 40°, e = 30 mpa; • mudstone, h = 2.1 m, e = 20 mpa. fig. 2 shows the geometrical dimensions of a metal corrugated pipe. in the pipe there was used corrugated sheet gl55 (fig. 3), which had the following characteristics: • thickness 4 mm; • the moment of inertia of the cross section i 181.92 cm4/m; figure 4. design scheme • the moment of resistance of section w 61.67 cm3/m; • the plasticity moment of resistance of section wp 79.72 cm3/m. the most common soil backfill for underground structures is sand. therefore, to evaluate the influence of soil characteristics on the stress-strain state of the shirred casing, two calculations were done. the first calculation was derived from a filling with sand with the following characteristics: e = 30 mpa; ϕ = 40°; c = 1 kpa. the second calculation was made for the filling out of fibrocartilagenous sand with the following characteristics: e = 30 mpa; ϕ = 45°; c = 67.0 kpa. the characteristics of sand reinforced with polypropylene fibres to an amount of 0.5 % is taken from test results given in [9–12]. according to the current design standard sp 35.13330.2011 “bridges and pipes”, the cargo load was a four-axle truck n14 with an axle load of 252 (kn), and the distance between the axles 1.2 m. the design scheme is shown in fig. 4. the results of calculating the stress-stain state of the tunnel lining with soil backfill are in figs. 5, 6, 7 and 8. analysis of the graphs showed that in the tunnel lining there occurs compressive strength with a maximum value of nmax = 147.87 kn and a considerable alternating bending moment with a maximum value 35 a. b. ponomarev, v.i. kleveko, o.v. moiseeva, k.r. kashapova acta polytechnica ctu proceedings figure 5. deformed scheme figure 6. the diagram of shell displacements for sand backfill (umax = 29.32 mm) of mmax = 6.5 knm. the maximum displacement of the tunnel lining was observed in the upper zone and amounted to 29.32 mm. the results from calculating the stress-strain state of the tunnel lining and soil backfill from fibroareolar sand is shown in figs. 9, 10 and 11. based on the calculation results the following data was obtained: • max. travel: imax = 20.71 mm; • max axial force nmax = 164.28 kn; • max bending moment mmax = 3.78 knm. the analysis of the above plots shows that using fibre reinforced soil instead of unreinforced soil leads to a decrease of the deformations of the tunnel shell by 8.6 %, a bending moment at 42 % and to an increase in longitudinal force of 9.7 %. based on the obtained calculation results there were calculated the normal stresses in the metal corrugated shell. the calculation results are presented in figs. 12 and 13. the analysis of these plots shows that the normal stresses in a metal enclosure when using the filling out fibre reinforced soil is much less than for soil that is unreinforced. figure 7. the diagram of axial force for sand backfill (nmax = 147.87 kn) figure 8. the diagram of bending moment for sand backfill (mmax = 6.5 knm) figure 9. the diagram of shell displacements for fibre reinforced sand backfill (umax = 20.71 mm) figure 10. the diagram of axial force for fibre reinforced sand backfill (nmax = 164.28 kn) 36 vol. 10/2017 fiber reinforced sand backfill figure 11. bending moment diagram for fibre reinforced sand backfill (mmax = 3.78 knm) figure 12. stress diagram for sand backfill (σmax = 105.39 mpa) 3. conclusions (1.) using fibre reinforced soil can improve the strength and deformation characteristics of the soil backfill corrugated metal tunnel lining, which leads to a significant reduction of stresses and strains in it. stress was reduced by 42 %, and the deformation by 8.6 %. (2.) the results of the calculations showed that improvement of the mechanical characteristics of the soil backfill for the tunnel allows the use of thinner corrugated sheets, which reduces the cost of construction. references [1] o. moiseeva, v. kleveko. analysis of accident cases involving pedestrians in perm. pnrpu construction and architecture bulletin (4):134–143, 2015. [2] p. a. tretyakova, v. kleveko. modern methods of increasing the efficiency of urban transport systems. pnrpu construction and architecture bulletin (1):101–108, 2012. figure 13. stress diagram for fibre reinforced sand backfill (σmax = 61.19 mpa) [3] o. v. moiseeva, v. i. kleveko. choice of a rational type of pedestrian crossing with safety for pedestrians and construction costs. world science and innovation 8(2):90–93, 2015. [4] a. e. polovnikova, v. i. kleveko. choice of a rational type of pedestrian crossings with regard to the safety of pedestrian traffic. modernization and research in the transport sector 2:356–361, 2015. [5] o. v. moiseeva, v. i. kleveko. choice of rational type of crosswalk with regard to the safety for pedestrians and the cost of construction. sworldjournal j21504.(2(9)):3–5, 2015. [6] o. v. moiseeva, v. i. kleveko. using prefabricated corrugated metal structures for the construction of a pedestrian tunnel. modern scientific researches and innovations 12(56):409–411, 2015. [7] v. v. kovalchuk. the effect of corrugated elements thickness on the deflected mode of corrugated metal structures. nauka ta progres transportu 3(57):199–207, 2015. [8] a. l. novodzinsky, v. i. kleveko. the account influence of a thickness of the goffered element on durability and stability of a metal water throughput pipe. pnrpu construction and architecture bulletin (1):81–94, 2012. [9] v. i. kleveko, y. m. shangina. the impact of the use of fibre reinforced sand as backfilling on the value of act. sworldjournal 04(116 (10)):16–21, 2016. [10] a. b. ponomarev, a. s. kuznetsova, v. g. ofrikhter. application of fibroreinforced sand as the foundation of buildings and constructions. bulletin of volgograd state university of architecture and civil engineering series: construction and architecture (30):101–107, 2013. [11] a. s. kuznetsova, v. g. ofrikhter, a. ponomarev. strength research of sand reinforced by discrete polypropylene fibres. pnrpu construction and architecture bulletin (1):44–55, 2012. [12] a. b. ponomarev, a. s. kuznetsova. laboratory testing of the strength characteristics of fibre reinforced sand with various water. bulletin of civil engineers 6(47):127–132, 2014. 37 acta polytechnica ctu proceedings 10:34–37, 2017 1 introduction 2 fibre reinforced sandy backfill 3 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0001 acta polytechnica ctu proceedings 11:1–5, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of language subset and language extension by safe related information approach to safe system development tomáš brandejský∗, vít fábera department of applied informatics in transportation, faculty of transportation sciences, ctu in prague ∗ corresponding author: brandejsky@fd.cvut.cz abstract. presented contribution is dedicated to discussion of two different approaches into increase of programming language safety. they are language subset and extension of original safety mechanisms. as examples we used misra c/c++ subset and spark language builded on the base of ada language. in the last chapters we discuss novel approaches based on application of programmable hardware which is described in vhdl language, which is also modification of ada language. especially spark and vhdl languages represents novel approaches to safe system development which are now discussed in relation to new railway sw projects. keywords: misra, ada, spark, vhdl, programmable hardware, software safety. 1. introduction nowadays we can observe two opposite approaches to modification of programming languages to be more suitable to safe system development. the first is the reduction of chosen programming language expressions to safe subset, like misra c/c++ [1, 2] and this approach is also supported by today standards of safe software development like iso61508 [3] or en50128 [4]. in the last few years, novel and partially opposite development is expanded. it is based on extension of original language by additional explanations containing description of such aspects of software, that are not described enough by original language. typical example is spark [5] language precisiating original ada one [6] even though it was originally designed as language specialized to safe control system development. 2. misra c/c++ misra c (motor industry software reliability association) c language subset was developed in 1998, latter it was updated and also there was developed analogical support for c++ (misra c++). misra c/c++ defines itself as guidelines. it defines language subset in the form of rules partially suggesting good practice solutions, frequently warning programmers away of wrong application of the language. especially valuable of misra guideline are parts explaining why was each rule formed, explaining what is the subject of possible future problem. applying of this rules allows to write robust code with decreased probability of malfunction. now it is possible to apply compilers and source code checkers implementing misra guidelines to allow automatic static analysis of code. even, there exists free software tool clang-misracpp2008 [7] but the development of this tool probably stopped before two years. from the practical viewpoint, misra guidelines checker implemented in compiler is also very interesting. it is even better than standalone checker because checking in compiler runs in each code making and does not require additional checking operation step. many companies producing compilers for development of embedded systems based especially on single chip micro-controllers, for example keil software company which is widely used by railway control system developers community in our country. there exists effort to implement misra rules into gnu c/c++ compilers because these compilers are free, proven in use and supports many processors and micro-controllers. one of the first works in this areas was [8], where authors describe its checker for gnu c which is written in ciao prolog language into which misra rules they were transformed. ciao prolog is similar to gnu prolog and is also free and compatible with gnu one but it seems to be more advanced than gnu prolog in its design. this project probably is not active now. there was also european project within the frame of itea3 call for development of gcc fork called ggcc [9] with aim to implement misra rule checking into gcc compiler. this project probably newer started or produced no reasonable result. thus it is possible to conclude that now there is no applicable support of misra for gnu c compilers and that gnu compilers development effort now focuses rather to support of parallel (super-)computers, like implementation of acc extension of standard c/c++ languages. misra c specification is not public domain and this fact might also cause small interest of gnu community to implement this subset. there exist also similar subset called high integrity c++ [10] by programming research, but also this specification is not public domain as well as checking tools. 1 http://dx.doi.org/10.14311/app.2017.11.0001 http://ojs.cvut.cz/ojs/index.php/app tomáš brandejský, vít fábera acta polytechnica ctu proceedings 3. ada language origin of ada language falls into falls into 80’s of previous century. this language was developed for us army applications as safe, strong typing programming language supporting parallelism, inter-process safe communication and defensive programming style preventing programmers errors and decreasing the need of testing. this pascal (or algol) like language was adopted in avionics and nuclear industry and it is referred by safe software system development standards [3, 4]. because the language was designed with respect to extremely large scale control system design operated by us ministry of defence, it was designed to eliminate maximum of programmers errors, ability to detect large percentage of them during compilation of code and with ability to operate in distributed environment with respect to possible communication errors and system component faults. the language was equipped by specific typing mechanism supporting defensive programming to support ability of error detection by compiler. defensive programming is related to strong typing, which means, that magnitudes between variables of different types cannot be shared without explicit conversion. because mainstream programming languages like c offer limited list of types (integer, real, character), the use of type control is able to discover only limited number of programmer errors when such languages are used. on the opposite side, ada language offers ways how to define new data types e.g. by constraining of definition domain of sample type (e.g. type secondfloornumbers is range 200..253; defines data type which magnitudes are constrained to really existing scale of numbers and allows to compiler to identify each impossible magnitude caused by programmer or communication error, e.g. 0). because strong typing might be exhausting, when it is required in situations when it brings to benefit, in ada it is possible to define subtypes, which magnitudes are transformed automatically without explicit type conversions between original type and all its subtypes. all data structures use closing clause unambiguously defining its end to prevent mistakes caused by wrong understanding of nested structures in the time of development, testing and maintenance. on the end of function and procedure definition it is supported its name stating. for many years, in mainstream programming languages the support of concurrent programming was extraordinary. ada language contains it from its origin in 70’s. ada allows not only define processes, but it contains strong support of safe inter-process communication and process synchronization called rendezvous. each process has defined so called entries to accept messages sent by other processes and is able to send messages to entries of other ones. but not only. it is possible to choose, which messages will be accepted in given state of process and it is possible to define reactions to unexpected situations, especially to fact, that given message was not received for given time. each process might define its local variables and separate itself from others. due to its real time orientation, ada language supports exception handling to solve unexpected situation without program fail. there is present concept of modular programming which applied templates (modules without precisely defined types, operators, procedures and functions which might be precisiated later in the time of final program linking. it allows e.g. to write sorting algorithm library without knowledge of sorted data type. such library si then applicable to any data type implementing assignment and comparison operators. latter ada adopted object oriented programming concept but it is not widely used as e.g. c/c++, java and some other languages, especially due to high price of commercial compilers and need of specific education of programmers. similarly like for c/c++, also for ada language there was effort to develop safe subset to increase its (high) safety. also standards [3, 4] refers this subset but in reality nothing such exists. the improving of ada language was moving different way and tends to spark language, especially due to impossibility do make such simple programmers errors as they are solved by misra for c/c++. significant part of such errors can not occur in standard ada and thus its improvement is much difficult, because even original (full scale) ada operates on different (much higher) level of safety than e.g. c. there exist support of ada language based on gnu compiler collection, but it means that there are supported especially versatile operating systems like linux and common purpose processors. there is also interesting, that compiler is not named ada compiler, but by acronym gnat. it is caused by licensing, when ada compiler term can be used only for compiler satisfying large (and thus expensive) verification, which is out of scope of gnu community. spark language is not pure subset. it also extends ada code by additional information to eliminate potential ambiguities and insecurities hidden in ada source codes. the main changes to standard ada language are the following: • the use of access types and allocators is not permitted (term access in ada denotes pointer – more frequent term known from the other programming languages, its misuse might cause significant errors hard to discover by testing). • all expressions (including function calls) are free of side-effects. this improvement was reached by incorporating of additional description of use of variables and symbols visible from neighborhood to prevent e.g. mistaking of global and local variable of the same name. • aliasing of names is not permitted in general but the renaming of entities is permitted as there is a static 2 vol. 11/2017 comparison of language subset and language extension relationship between the two names. in analysis all names introduced by a renaming declaration are replaced by the name of the renamed entity. this replacement is applied recursively when there are multiple renames of an entity. • the goto statement is not permitted. • the use of controlled types is not currently permitted. • tasks and protected objects are permitted only if the ravenscar profile (or the extended ravenscar profile) is specified. ravenscar profile denotes set of pragmas influencing compiler work to prevent some possibly unsafe construction in ada code. ravenscar profile is available in standard ada too, in spark is is not option but he standard. • raising and handling of exceptions is not currently permitted (exceptions can be included in a program but proof must be used to show that they cannot be raised). misra c/c++ especially eliminates confusing structures which might be understand differently by compiler, programmer and structures which are differently understand by different compilers, especially due to vague language definition (it is especially problem of c language, c++ in its last version was significantly improved, precisiated). for example if there is used condition ((x==a)&&(y==b)), there is not defined in language specification if the sub-expression (x==a) is evaluated as the first or the second. it is no problem until we do modification of this condition to form ((x()==a)&&(y()==b)). in this new condition x and y are not static objects (variables), but they are function calls and functions might contain so-called side effects, in this case they might operate the same global variable and thus the results of calls of functions x() and y() might be influenced by call order. this subset has also many rules eliminating confusing structures causing misunderstandings by programmers and testers. misra subsets contain tens such confusing structures. e.g. it is possible to mention constants starting with zero (they are understood by compilers as octal numbers by definition), problem of nested structures if-else which might not contain else branch and no terminal symbol stopping this structure is defined in the language and so on. 4. spark spark language introduces explicit description of use of global objects (and commonly objects defined on higher level of visibility) called dependencies. the use of these additional descriptions is known from design by contract methodology, which was on the origin of such languages as eiffel [11]. this methodology prescribes that software designers should define formal, precise and verifiable interface specifications for software components, which extend the ordinary definition of abstract data types with preconditions, postconditions and invariants. preconditions are conditions which must be satisfied to enter the following section. opposite postcondition must be satisfied on the end of the block and the invariant must be valid inside the block. 5. vhdl vhdl (vhsic hardware description language) [12] is now frequently used in programming of field-programmable gate arrays. it describes structure and/or behaviour of digital hardware. vhdl is similar to ada language both in concept and syntax (parallel processes, subtypes, ...). it was primarily proposed for simulation, now the vhdl description is an input for synthesis tool instead of schematic diagram to produce final design. the interesting question is, if the similarity of vhdl and ada languages cannot be used to description of both sw a fpga implementation of systems with high requirements to reliability, applicability, maintenance and safety. simulation of system described by vhdl language is not straightforward due to need to respect signal and not process oriented nature of fpga. this nature caused presence of sensitivity list in vhdl code. in vhdl, the proces statement has an part called sensitivity list. the syntax is as follows: [process_label:]process[(sensitivity_list)] declarative_part begin statement_part end process[process_label]; example: andgate1: process (a,b) begin y <= a and b after 5 ns; end process; the sensitivity list consists of input signal identifiers separated by commas. sensitivity list is important at the moment of the simulation. during the initialization phase of simulation, the process runs and it assigned initial values of signals and variables (signal y in our case). then the process is suspended. if the any signal listed in the sensitivity list is changed, the process is resumed and executed. the sensivity list after the process keyword is syntactically optional, but we can really say that is mandatory. the sensitivity list placed after the process keyword can moved at the end ot the process section and written like the wait on statement: andgate1: process begin y <= a and b after 5 ns; wait on a, b; end process; 3 tomáš brandejský, vít fábera acta polytechnica ctu proceedings figure 1. corresponding schematic diagrams to vhdl code related to simulation. the process having sensitivity list after process keyword must contain no wait on statement. if some signal is missing in the list, it affects naturaly the simulation. for example, signal b is missing: andgate1: process (a) begin y <= a and b after 5 ns; end process; let consider the set of simulus: both signals are set to 0 at time 0ns, a signal is changed to 1 at 10ns, b signal is changed to 1 at 20 ns. the result of simulation in isim (a part of xilinx ise design pack) is constant zero signal y, because the process in not activated at the time 20 ns due to missing signal b in the sensitivity list. while process andgate1 really describes combinational and, andgate2 process has sequential behavior. we show one more example of code and corresponding schematic diagrams to simulation (fig. 1). multiplexor1: process (s, i0, i1) begin if s = ’0’ then y <= i0; else y <= i1; end if; end process; multiplexor2: process (s) begin if s = ’0’ then y <= i0; else y <= i1; end if; end process; in simulation, the output y of multiplexor2 process changes only if event on s occurs. it means that output y follows the value of i0 at the moment of falling edge on s and value of i1 at the moment of rising edge on s. 6. synthesis processes andgate2 and multiplexor2 have sequential behaviour during the simulation with isim. how process these codes synthesis tools xilinx ise 12.4? if a signal is missing in the sensitivity list, tools writes a warning like "i0 should be on the sensitivity list of the process". if signal y is listed in the sensivity and it y an output of the entity, the vhdl compiler writes an error "y with mode ’out’ cannot be read". if y is an internal signal, it can state in the sensitivity list without warning or error. but all variants of code lead to the same rtl (register transfer level) – combinational logic – and gate and multiplexor. syntesis tool decides if it uses combinational logic or flip-flop according to typical vhdl code structures (if clk’event and clk = ’1’ then for flip-flops). so, if we forget to put a signal in the sensitivity list and if we ignore warnings, the simulation of the code and the behavioral of the target design can be different. the reason is: to put all input signal in the sensitivity list and or use a variant of the sensitivity list with keyword all, which was added into new specification of the language. the gcc compiler has -wall switch, that cause to stop compiling also if a warning occurs. switch -wall is recomended in c language if we create a safety software. synthesis tools (xilinx ise) have usually no switch like -wall, vivaldo design suite from xilinx makes possible to change severity from warning to errors. as seen, vhdl, although declared as language with safety properties, has also some weak points. a question is: wouldn’t it be suitable to extend vdhl compilers in the sense of -wall switch and to add a variable list like in sparc, espetially, if vhdl is speculated as a language for railway systems? by the way, rules for safety design using vhdl was presented in [13]. 7. discussion vhdl language uses control structures and programming methodology analogical to original ada language. it is the subject of future research, if it is possible to modify it in the way analogical to spark to increase its safety and analysis if such extension could bring significant increase to safety due to influence of hardware (gate array) to it. 8. conclusion the presented paper brings analysis of impact of two basic methods to increase language safety to cover safety requirements of systems working on high safety integrity levels, e.g. on railways. it is not only focused to comparison of ada and spark languages but it is illustrated on problems bring by implementation of vhdl simulator designed with respect to describe safe system by single language without need to distinguish if this system will be implement by hw or sw. references [1] guidelines for the use of the c language in critical systems, march 2013. isbn 978-1-906400-10-1 (paperback), isbn 978-1-906400-11-8 (pdf). 4 vol. 11/2017 comparison of language subset and language extension [2] guidelines for the use of the c++ language in critical systems, june 2008. isbn 978-906400-03-3 (paperback), isbn 978-906400-04-0 (pdf). [3] iec 61508 – electronic functional safety package. [4] en50128: railway applications – communications, signaling and processing systems – software for railway control and protection systems, 2011. [5] iso/iec 8652:2012(e). ada reference manual. copyright ©1992, 1993, 1994, 1995. [6] spark 2014 reference manual. copyright ©2013-2016. [7] "https://github.com/rettichschnidi/clangmisracpp2008". [8] g. marpons. a coding rule conformance checker integrated into gcc [online]. "babel.ls.fi.upm.es/ gmarpons/pubs/prole08codingrules.pdf". [9] ggcc reseny mandrivou v ramci itea3. [10] "http://www.codingstandard.com/section/index/". [11] eiffel language [online]. "http://www.ecmainternational.org/publications/standards/ecma367.htm". [12] ieee standard vhdl language reference manual, 2000. [13] t.musil. návrh metodiky pro vývoj a verifikaci bezpečných algoritmů implementovaných v dynamicky rekonfigurovatelných fpga (methodology for design and verification of algorithms with demands for safety on dynamically reconfigurable fpga), doctoral thesis. czech technical university, 2015. 5 acta polytechnica ctu proceedings 11:1–5, 2017 1 introduction 2 misra c/c++ 3 ada language 4 spark 5 vhdl 6 synthesis 7 discussion 8 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.10.0007 acta polytechnica ctu proceedings 10:7–11, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app correlation relationships for the hypoplastic model for fine grained soils tomáš kadlíček∗, tomáš janda, michal šejnoha department of mechanics, faculty of civil engineering, czech technical university in prague, thákurova 7/2077, 166 29 prague, czech republic ∗ corresponding author: tomas.kadlicek@fsv.cvut.cz abstract. the paper is concerned with our ongoing research effort devoted to the development of reliable computational tools for the calibration of advanced constitutive models of soils. at present, such software is available for the hypoplastic model of clays applicable to soft soils. this software provides a stepping stone for the determination of potential links between individual model parameters and fundamental characteristics of soils. identifying such links would allow for tuning the model without performing time consuming experiments, particularly in the case of an initial design. some preliminary results are presented in the paper. keywords: hypoplasticity, fine grained soils, constitutive modeling, correlation. 1. introduction the theory of hypoplasticity has been receiving considerable attention in recent years, particularly for its capability of well simulating the basic characteristics of soils resulting in reliable predictions at a structural level. these characteristics include not only the dependency of stiffness on the void ratio, but also the stiffness dependency on strain and the loading history. such relationships are hardly available with frequently used elastoplastic models of the mohr-coulomb (mc). these are generally adopted in fem simulations because of their simplicity and straightforward links to basic strength parameters such cohesion and the angle of internal friction, which are included in regular laboratory test protocols or available in common material databases. to provide a similar comfort or at least considerably simplify the use of hypoplastic models is the principal objective of our current research focusing particularly on the masin hypoplastic model for fine grained soils (hc) [1] and the von wolffersdorff hypoplastic model for coarse grained soils (vw) [2]. in particular, two specific directions are investigated. the first aims at creating easy to use calibration software built upon standard oedometric and triaxial laboratory tests. the second option assumes a calibration procedure in the form of tables or graphs enabling the determination of model parameters on the basis of specific input data. knowing the link between the model parameters and more standard soil parameters creates the basis for the development of such a tool. this objective is examined in this paper and some preliminary results are presented. 2. hypoplastic model for fine grained soils the theory of hypoplasticity can be considered a possible alternative to the theory of elastoplasticity. unlike elastoplasticity, it does not separate an elastic and plastic deformation, but its deformation is controlled by the evolution of state variables, namely pressure and the void ratio. consequently, hypoplasticity exhibits a distinct nonlinear deformation both in primary loading and elastic unloading. the state variables are bound in the e×q×p (e âăş void ratio, q deviatoric stress, p mean stress) space, where all admissible states are enclosed by the so called state boundary surface (sbs). the hc model defines five material parameters: λ∗, κ∗, n, ϕc and ν. these parameters can be calibrated on the basis of simple oedometric and triaxial laboratory tests. when properly tuned the model provides reliable and accurate predictions of the soil response in the solution of real engineering tasks. the first three parameters: λ∗, κ∗ and n are closely related to the oedometric test. as seen in fig. 1, the parameter λ∗ controls the slope of the normal consolidation line (ncl). the position of ncl in the ln p× ln(e + 1) space is driven by the parameter n. clearly, recall fig. 1, its value corresponds to the mean pressure p = 1 kpa. the third parameter κ∗ represents the slope of the swelling line followed upon unloading as also evident from fig. 1. for more details, the interested reader is referred to [1] and [2]. the remaining two parameters ϕc and î¡ are typically determined on the basis of the triaxial test. in the case of fine grained soils, an undrained triaxial test accompanied by measuring the pore pressure is preferred. the critical state friction angle ϕc is dis7 http://dx.doi.org/10.14311/app.2017.10.0007 http://ojs.cvut.cz/ojs/index.php/app t. kadlíček, t. janda, m. šejnoha acta polytechnica ctu proceedings figure 1. parameters λ∗, κ∗ and n in the ln p × ln(e + 1) space figure 2. critical friction angle ϕc in the t′ × s′ space played in fig. 2. the effective stresses t′ and s′ in fig. 2 are written as s′ = σ′1 + σ′3 2 t′ = σ′1 −σ′3 3 (1) the last parameter î¡ is related to the ki/gi ratio, where ki is the bulk modulus in isotropic compression and gi is the shear modulus in undrained shear for a test starting from the isotropic normally compressed state. obtaining a reliable prediction of this parameter requires a parametric study. the result of such a parametric study in terms of an equivalent deviatoric stress plotted as a function of the vertical strain is shown in fig. 3. additional details are available in [1] and [2]. further details on the proposed calibration procedure for both hc and vw models, developed as well as validated on the basis of a large ensemble of laboratory tests, are provided in [3]. in the next section, limiting our attention to the hc model, the model parameters determined for a representative set of soil samples will be correlated with a selected set of other, easily determined, soil characteristics. 3. hc parameters’ correlations soils can be characterized and classified according to the particle-size distribution and atterberg limits. in order to determine the size distribution of particles in a specimen the sedimentation test (for grain sizes < 0.063 mm) and measurements of the amount of particles passing a particular sieve size can be performed to derive the grain size distribution curve. the atterberg limits are determined as the values of mass moisture content of a soft soil and are defined as follows: • shrinkage limit ws is the water content where an ongoing loss of moisture is not followed by any additional volume reduction • plastic limit wp is the water content where cylindrical samples of soil being 3mm in diameter fall apart in pieces 8-10mm long • liquid limit wl is the water content where a groove made in soil placed in the casagrande device merges at a length of 10mm after 20 drops. in the present study, the results from oedometric and triaxial measurements and index tests carried out on eight samples of soft soils were exploited. while the first set of experiments was employed in the calibration procedure to assess the values of individual model parameters, the actual soil classification was based on the index tests. the resulting values are listed in tab. 1. apart from model parameters and the atterberg limits wl and wp, the table also offers the percentage content of fine soil particles fwith the diameter d < 0.063mm. the particle-size distribution is also plotted in fig. 4 for the sake of clarity. when searching for possible correlations between the model parameters and the selected soil characteristics it has been proven convenient to adopt parameters calibrated on the basis of reconstituted soil samples. these samples were prepared from the original undisturbed soil by changing the sample paste consistency up to a water content of 1.5 × wl while completely disrupting their original internal structure and as such losing any information about the prior loading history. the values of model parameters are nevertheless the same as those which would have been found when adopting the original undisturbed samples. the relationships between the values of hc parameters and the index characteristics, see fig. 5 and 6, can be conveniently specified in terms of the coefficient of correlation r ∈ (−1; 1). recall that the correlation reflects the noisiness and direction of a linear relationship of two random variables, but not the slope of that relationship. when there is no obvious relationship 8 vol. 10/2017 correlation relationships for the hypoplastic model figure 3. evolution of the equivalent deviatoric stress q as a function of the vertical strain εa for various values of ν table 1. parameters of the hc model and selected parameters from index tests for reconstituted soil samples figure 4. particle-size distribution for eight samples of sampled soils 9 t. kadlíček, t. janda, m. šejnoha acta polytechnica ctu proceedings figure 5. correlations between hc parameters and index characteristics (correlation), the coefficient of correlation is equal to 0. in the case of a perfect direct (increasing) correlation, r approaches 1 and in the case of a perfect inverse (decreasing) correlation, r approaches -1. thus, the closer the coefficient is either to -1 or 1, the stronger the correlation between the two variables. tab. 2 shows the coefficient of correlation for all five hc model parameters and the selected index characteristics wl and wp, and the percentage content of fine soil particles f. a relatively strong (direct) correlation is observed for the model parameters λ∗ and n and the content of fine soil particles f. this is also confirmed by the plots in fig. 5(a,b) showing no apparent scattering. remember that the parameters λ∗ and n control the position of ncl in the ln p× ln(e+1) space. both tab. 2 and the plots in fig. 5(c,d) suggest the parameter κ∗ be more dependent on the plasticity index wp and the liquid limit wl. the same conclusion, although not directly visible from tab. 2, can be drawn for the parameter v, see fig. 5(e,f). for this parameter, the value of the coefficient of correlation is polluted by the results derived from the soil sample labeled as “praha a”, which falls off the more or less linear relationship. this specimen 10 vol. 10/2017 correlation relationships for the hypoplastic model table 2. correlation values r figure 6. correlations between hc parameters and index characteristics shows the lowest values of the atterberg limits and the lowest content of fine soil particles f. unlike the first four parameters associated more or less with the compressive response of soil, the critical state friction angle ϕcv does not seem to show any correlation with the selected index characteristics, see both tab. 2 and fig. 6(g). however, its determination is rather simple from a standard triaxial test with no need for additional optimization. 4. conclusion although drawing a solid conclusion as to the direct link of the hc model parameters to other, more standard, soil parameters would require a considerably larger set of experiments, the potential of this goal has been clearly confirmed by this study. in particular, a relatively strong correlation of the hc model parameters λ∗, κ∗, n and ν and the selected index parameters or the particle size distribution data can be found. while limiting our attention to the hc model we expect that a similar study performed for the vw model can lead to similar conclusions. obviously, this must be supported by a much larger set of data to be examined. we expect such information to grow with the help of a web application which not only provides calibration software, but opens a way to collected results from laboratory measurements all over the world. [4] 5. acknowledgement the financial support from projects ta04031603 and sgs16/038/ohk1/1t/11 is gratefully acknowledged. references [1] d. mašín. clay hypoplasticity with explicitly defined asymptotic states. acta geotechnica 8(5):481–496, 2013. doi:10.1007/s11440-012-0199-y. [2] d. mašín. a hypoplastic constitutive model for clays. international journal for numerical and analytical methods in geomechanics 29(4):311–336, 2005. doi:10.1002/nag.416. [3] m. šejnoha, t. kadlíček, t. janda. calibration of hypoplastic models for soils. in engineering mechanics 2015, vol. 821 of applied mechanics and materials, pp. 503–511. trans tech publications, 2016. doi:10.4028/www.scientific.net/amm.821.503. [4] p. a. von wolffersdorff. a hypoplastic relation for granular materials with a predefined limit state surface. mechanics of cohesive-frictional materials 1(3):251–271, 1996. doi:10.1002/(sici)10991484(199607)1:3<251::aid-cfm13>3.0.co;2-3. 11 http://dx.doi.org/10.1007/s11440-012-0199-y http://dx.doi.org/10.1002/nag.416 http://dx.doi.org/10.4028/www.scientific.net/amm.821.503 http://dx.doi.org/10.1002/(sici)1099-1484(199607)1:3<251::aid-cfm13>3.0.co;2-3 http://dx.doi.org/10.1002/(sici)1099-1484(199607)1:3<251::aid-cfm13>3.0.co;2-3 acta polytechnica ctu proceedings 10:7–11, 2017 1 introduction 2 hypoplastic model for fine grained soils 3 hc parameters' correlations 4 conclusion 5 acknowledgement references acta polytechnica ctu proceedings doi:10.14311/app.2016.6.0018 acta polytechnica ctu proceedings 6:18–27, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app framework for ad hoc network communication in multi-robot systems khilda slyusar∗, miroslav kulich czech technical university in prague, zikova 1903/4, 166 36 prague, czech republic ∗ corresponding author: slyuskhi@fel.cvut.cz abstract. assume a team of mobile robots operating in environments where no communication infrastructure like routers or access points is available. the robots have to create a mobile ad hoc network, in that case, it provides communication on peer-to-peer basis. the paper gives an overview of existing solutions how to route messages in such ad hoc networks between robots that are not directly connected and introduces a design of a software framework for realization of such communication. feasibility of the proposed framework is shown on the example of distributed multi-robot exploration of an a priori unknown environment. testing of developed functionality in an exploration scenario is based on results of several experiments with various input conditions of the exploration process and various sizes of a team and is described herein. keywords: multi-robot systems, distributed exploration, ad hoc network, limited communication, cooperation. 1. introduction one of the key issues when designing a robotic system involving more than one entity is how these entities communicate. the majority of current multi-robot solutions are centralized (i.e., they have a central element through which all messages between robots are passed) or assume an unlimited communication range allowing to send messages directly between two arbitrary robots. this setup has many advantages, but also several drawbacks. it can be hard, costly or even impossible to prepare external communication infrastructure assuring one-to-one communication accessibility in hostile, unexplored or large environments. moreover, robustness of centralized systems which are highly dependent on a single node is low. finally, it might be difficult to exchange information in large systems consisting of many robots as their communication capacities are limited and a number of messages is high. some other arguments as well as real-world applications where ad hoc networks improved or might improve performance, scalability and robustness of robotic teams can be found in [1]. elkady et al. [2] give a nice survey of robotic middlewares. besides centralized client-server architecture employed, e.g. in player/stage/gazebo [3], ace/tao library [4], implementation of corba (common object request broker architecture) standard [5], is used for peer-to-peer communication in middlewares like claraty [6], miro [7], orocos [8]. alternatively, the ice framework by zeroc [9] for remote procedure calling (rpc) can be used as it is already done in the orca project [10]. both zeroc ice and corba do not provide communication between directly inaccessible nodes in a decentralized fashion and although extensions for ad hoc networks exist [11] they were not used in the mentioned robotic middlewares. communication in robot operating system (ros) [12], the mostly popular robotic framework, is peer-to-peer, but this communication is mediated by so called rosmaster. this node provides naming and registration services and enables individual ros nodes to locate one another. once the nodes locate each other, they communicate directly. also ros does not provide routing of messages between inaccessible nodes. on the other side, several works were presented which describe application of mobile ad hoc networks with routing in robotics. witkowski et al. [13] present a cell-based network with master nodes in each cell and routing between distant nodes is provided through these master nodes which are distributed uniformly in the environment. the authors validate the approach on two scenarios. while measuring of the signal quality of wlan and bluetooth communication nodes in indoor environments with different kind of wall materials is the aims of the first one, the second experiment deals with radio-based positioning of robots in a dynamic environment. corell et al. [14] study a wireless coverage problem with minimal requirements to hardware of robots. as a communication middleware they use optimized link state routing (olsr) [15] which provides tcp/ip routing for the resulting mesh network. a suite of algorithms that cover a defined area by a fully distributed robotic team, detect network disconnections, perform network repair, and repair area coverage is presented in [16]. finally, agrawal et. al [17] experimentally evaluate two routing protocols for a robotic swarm. the swarm aims to find a source of heat and a particle swarm optimization algorithm is employed as a search strategy. the aim of this paper is to give an overview of routing protocols in ad hoc networks and choose the most 18 http://dx.doi.org/10.14311/app.2016.6.0018 http://ojs.cvut.cz/ojs/index.php/app vol. 6/2016 framework for ad hoc network communication in multi-robot systems suitable for application in the multi-robot domain. the chosen protocol is than used in a design and realization of the framework for exploration of an a priori unknown environment which thus serves as a testbed for experimental evaluation of the protocol implementation. the rest of the paper is structured as follows. section 2 contains a structured list of requirements for a routing protocol, a classification of protocols, and describes the main representatives of the categories. then a comparative analysis of routing protocols is performed. analysis of three message-oriented middlewares (mom) describes in section 3. the general application architecture and using of selected measures is presented in section 4. section 5 contains results of the experiments, which allow to verify the implementation performance of the proposed architecture. the final evaluation of the paper is in conclusion. 2. ad hoc routing protocols a robot team forms an ad hoc network as it does not rely on a pre-existing infrastructure, while interaction between team members is done by a routing protocol. operation efficient of a distributed robotic system requests a routing protocol to meet certain requirements. the key assumption is fully distributed operations and the protocol should not depend on a central control node. due to high mobility of the nodes, the protocol should easily adapt to changes in network topology caused by movement of nodes. it also needs to be localized, since global exchange of routing information will require large overheads [18]. an important requirement is the absence of routing loops. a routing loop occurs when the packets continue to be routed in an endless circle of the nodes. the protocol should not contain them, otherwise the overall efficiency is reduced, the routing loop will spend more bandwidth and more computing power of each node in this loop. the protocol should take care about saving resources of network members. in addition, the protocol needs to be scalable, this means that its effectiveness should not depend on the number of members in the network. increase or decrease of the network should not affect its performance. it also needs to be adapted to any speed of nodes and any type of their movement. in conditions of high mobility the protocol must be able to quickly recalculate cost and build the path to the destination which will include the lowest number of other nodes. while building the path and forwarding of messages, it is necessary to effectively avoid outdated routes [19]. the rest of the section contains a brief description of few most commonly used routing protocols and their comparison based on the above requirements. the main purpose of this comparison is to select the routing protocol that is the most suitable for teams of mobile robots which establish a wireless mobile ad hoc network (manet) without any centralized structure. the routing protocols can be divided into proactive (table-driven), reactive (on-demand) and hybrid, which combines advantages of the first two categories. 2.1. proactive routing protocols table-driven protocols constantly maintain the relevance of the routes between sources and destinations. to maintain a correct view of the network topology a protocol of this type responds to every change in the structure by sending changes across the entire network. moreover, the protocol periodically sends route update messages to its neighbours. these protocols require to store routing information in one or more tables in each node [20]. these protocols do not scale well and control overheads are proportional to the number of nodes in the network. destination sequenced distance vector (dsdv) [21] is one of table-driven protocols and it is a loop-free modification of the bellmanford routing algorithm [22]. each node maintains a routing table that contains entries for the next hop on the shortest path to all reachable destinations. the protocol uses a sequence number attribute to ensure the freedom of cycles and to distinguish stale routes. dsdv uses two types of dumps with updates of the routing table which sends to all immediate neighbours. the first type is periodically sent full dump which contains all available information about the routing information. the second type is an immediately sent incremental dump which sends on route changes and contains only changes that have occurred since the last full dump. another proactive protocol is optimized link state routing (olsr) [15] protocol, which is based on basic link-state algorithm. every node maintains a network topology graph. each node has thus shorter routes to every destination immediately, when data transmission begins. to keep the relevance of the topology a node periodically floods hello messages about its available links to the others. olsr uses also topology control (tc) messages that contain information about one-hop neighbours. the main optimization lies in using of multipoint relays (mprs). it reduces message overhead, because only nodes which were chosen as mprs forward broadcast messages during the flooding process. 2.2. reactive routing protocols reactive protocols perform the process of route discovery only on request. the source floods the network with route query requests when a packet needs to be routed using distance vector routing or source routing. the problem of reactive protocols is the delay of packet transmission during the route discovery. on the other side, the route is discovered only when needed, i.e., it is generally less memory demanding compared to 19 khilda slyusar, miroslav kulich acta polytechnica ctu proceedings proactive protocols and requires relatively little control traffic overhead. dynamic source routing (dsr) [23] is an example of on-demand routing protocols. dsr supports two types of operations: route discovery, when the node is required to transfer the data to the destination with an unknown route, and route maintenance, that allows to determine that the network topology is changed. after the work of these mechanisms the sender knows the complete hop-by-hop route to the destination. all transmitted routes stored in a route cache decrease time of route discovery. the protocol allows multiple alternative routes to the same destination and allows each node to perform load balancing. another example of reactive protocols is temporally ordered routing algorithm (tora) [24] based on the link-reversal algorithm. tora performs three functions: route creation, when a node uses a height metric to establish a destination oriented directed acyclic graph, route maintenance, when a node reestablishes routes due to topology changes and route erasure, when it is necessary to erase invalid routes. the protocol builds loop-free routes and provides multiple routes to alleviate congestion. however, tora may produce temporary invalid routes, as well as the light-weight mobile routing protocol [25] which it is based on. 2.3. hybrid routing protocols hybrid protocols typically try to reduce delay of route discovery from reactive systems by creating some form of routing tables and reduce control traffic overhead from proactive systems [20]. zone routing protocol (zrp) [26] is a hybrid protocol, which uses a proactive approach to a close neighborhood and a reactive approach to remotely lying nodes of the network. in other words, each node divides the network into intrazone and interzone. the node immediately knows routes to all nodes within the intrazone. the discovery of such paths is produced by the algorithm similar to dsdv. if an initiator needs to send data to a recipient which is in the interzone the sender proactively maintains a route to the destination. this approach reduces latency in route discovery and decreases the number of control messages. ad hoc on-demand distance vector routing (aodv) [27] is also a hybrid protocol based on the principles of the dsdv and dsr protocols. the protocol sends a periodic hello message, maintains the route table with a maximal single route for each destination and uses the principle of sequence numbering, as dsdv routing protocol does. on the other side, aodv provides a similar route discovery procedure as in dsr. a route between nodes is established only if it is necessary to send the data and there is no active route to the destination, i.e., the routing protocol reduces network overhead as compared with the dsdv algorithm. moreover, the route discovery packets do not contain a complete sequence of nodes, that is an advantage compared to dsr. 2.4. comparison of routing protocols there are several criteria [28] for evaluating the performance of a routing protocol: • packet delivery fraction is calculated as the ratio between the number of received messages by the destination to the number of messages sent by the sender; • end-to-end delay is the average time that elapses between the first packet which was transmitted by the sender before the first data packet received by the destination. the metric includes the time of route discovery, transmission delay, queuing delay and propagation delay; • routing overhead is calculated as the ratio between the routing packets to the total number of packets transmitted by the sender; • throughput is the amount of successfully delivered data via a communication link per time. it is usually measured in bits per second. the article [28] discusses three routing protocols dsdv, aodv and dsr. testing is carried out on teams of robots, whose size ranges from 10 to 80 pieces. robots perform a task similar to the task of terrain exploration. each node chooses some point in the terrain and moves toward it. the article considers all four criteria. dsr and aodv show better value of packet delivery fraction than dsdv. the value is about 95% for a team of 10 robots and it tends to 100% with an increase in the team size. aodv shows the smallest value of end-to-end delay. dsdv is a proactive routing protocol, so it shows better results than dsr. the value of routing overhead for aodv is better than dsdv. however, dsr has the smallest routing overhead in comparison with the other two. throughput has the lowest value for proactive dsdv. both dsr and aodv show approximately the same results, but this metric for dsr is slightly higher. a performance comparison of aodv, olsr and zrp is described in [29]. the simulation is conducted for teams of robots with sizes of 25, 50, 75, 100, and provides measurement of all metrics, except the routing overhead. the packet delivery fraction for aodv is approximately 90%, while olsr and zrp have the values of 20-40%. moreover, aodv has the highest throughput, as compared to two other routing protocols. however, aodv shows a significantly larger end-to-end delay than the reactive olsr and the hybrid zrp. the publication [30] compares the performance of dsdv, tora, dsr and aodv. the maximum number of robots in the team in the simulation is 50. the article considers several metrics, which include packet delivery fraction. both dsr and aodv show excellent results for the packet delivery ratio, which 20 vol. 6/2016 framework for ad hoc network communication in multi-robot systems activemq apollo zeromq nanomsg open-source yes yes yes has detailed documentation yes yes only manual pages has new updates yes yes yes point-to-point message passing yes yes yes publish/subscribe message passing yes yes yes written in java c++ c transport protocol tcp, udp tcp, in-proc tcp, in-proc support brokerless model no yes yes complexity of use requires broker without without configuration pre configuration pre configuration table 1. a brief comparison of message oriented middlewares. ranged from 95 to 100% depending on the network load. aodv is selected for implementation of the exploration framework on the basis of the above mentioned comparisons. simulations use robot teams of various sizes and display measurement results that are close to the conditions of the real world. the main advantage of aodv is a high percentage of packet delivery to the destination. the routing protocol has a good throughput compared to the rest. the designed application does not send many unicast messages, so the inflated end-to-end delay and routing overhead are not important. that is, these values are acceptable for the terrain exploration problem. 3. message oriented middleware the exploration application uses message passing to communicate between team members. it is a technique for invoking behavior, which uses incoming messages from other processes to run a code. there are two types of message passing: synchronous and asynchronous. the synchronous approach implies that message exchange takes place at a time when applications are running simultaneously. it uses temporary sender blocking before receiving a response from a recipient and infinitely blocking in case of recipient unavailability. the robots move in the environment during the exploration problem solving and can get out of the communication range. in addition, the robot can be turned off, for example, due to the battery discharge. in such cases, the message sender is blocked. that is, synchronous approaches are not suitable for use in distributed robotic systems that solve exploration problem. the asynchronous approach does not require simultaneous operation of a sender and a recipient. the intermediate level of software provides all the operations for receiving, storing and sending of messages. the most known type of such software is messageoriented middleware (mom). table 1 provides a brief comparison of main parameters of these three widely used asynchronous message-oriented middlewares: activemq apollo [31], zeromq [32] and nanomsg [33]. as can be seen from the table, all three libraries are open-source and release new updates. the first two libraries are well documented and have a large community of users. the latter is improved and rewritten zeromq library, which currently has only the manual pages, but the number of nanomsg users is constantly growing. all considered libraries support binding for commonly used programming languages as java, c, c# and c++. zeromq and nanomsg have also bindings for ruby, go, haskell, perl, php and others. there are two main functionalities of moms: message queuing and publish/subscribe. the first approach is a point-to-point messaging model, that is, a message is sent from a sender to a recipient. publish/subscribe approach is a many-to-many messaging model. all three libraries include both types of message passing. the libraries support a wide range of transport protocols, but transmission control protocol (tcp), user datagram protocol (udp) and in-proc are relevant for the purposes of the application. all considered libraries support transfer of data over tcp. activemq apollo additionally supports udp, while the other two do not support it. zeromq and nanomsg can use in-proc transport protocol. it is an inter-thread transport, which is much faster than tcp and can transfer data between threads of a single application. the most important difference is that the basic principle of activemq apollo is to create and configure a broker. a distributed multi-robot system has no central node, which could be a broker. a possible solution is creation of a broker at each node of the network. however, this solution is cumbersome and requires excessive pre-configuration of brokers. contrariwise, 21 khilda slyusar, miroslav kulich acta polytechnica ctu proceedings zeromq and nanomsg support brokerless model. nanomsg library was developed by one of the key creators of the zeromq library and it fixes most of the zeromq drawbacks. the main one is the problem with fault tolerance. nanomsg provides a new pattern of communication that does not require redundant confirmation from a recipient in the model requestrouter that resembles the work of procedure calls. the library allows to set a timeout for an operation of sending messages. it allows to refuse the message sending, if it is not sent for some predetermined amount of time. on the contrary, zeromq does not provide the possibility to obtain information about the current status of the recipient before sending a message. moreover, the library does not allow to set a timeout on an attempt to send a message. thus, in case of a fall, the whole system of robots gradually stops because of waiting a response from each other. the nanomsg library was chosen for use in the application on the basis of the above comparison of the characteristics of the libraries. because the library doesn’t provide an implementation of udp, the transport protocol is proposed to implement using standard sockets. 4. distributed exploration exploration as a process of autonomous navigation of a mobile robot in order to build a model of an a priori unknown working environment in a shortest possible time is one of the fundamental problems in mobile robotics. in general, exploration is an iterative process that performs the following operations at each step: the robot receives information from its sensors, updates an internal model of the surrounding space on the basis of these data, selects the next goal for the navigation on the basis of the current knowledge, and gradually moves to the selected destination. the process continues until there are no more available goals, that is, all areas of the environment will be explored [34, 35]. the main source of optimization is preparation of a set of possible goals and selection of a next goal from them, which is the most favorable for exploration purposes. besides the tasks mentioned above, the problem of distributed terrain exploration by a multi-robot team requires solving the problem of communication between team members and the problem of coordination their actions. this is why we selected this task as a testbed for evaluation of the aodv routing protocol. information about the terrain which is exchanged between team members and the method of determining the set of possible navigation goals depends on an exploration approach that is selected for implementation. frontier-based exploration [34] is chosen as one of the most common and widely used approaches which also has the extension for the case of multi-robot exploration [36]. the approach is based on the concept of a frontier, which is a boundary between the already known part of the terrain and still an unknown part. the internal representation of the map is an occupancy grid, which uses bayesian based updating technique. the main advantage of the environment exploration by a multi-robot team is the ability to easily share knowledge about the terrain between its members. exploration of non-overlapping areas by different robots allows to raise efficiency of teamwork and reduce the total exploration time. each robot in the team conducts the current goal selection independently on the basis of its current knowledge of the environment and the goals of the other team members. a greedy algorithm is chosen as the most primitive and native approach. it means, that the closest frontier to the current position of the robot is selected as the goal. in addition, the selected goal should not lie in the areas of the current goals of the other team members, which have active communication link to the robot. after the goal selection the robot informs the others about the selected goal and reports the priority for this goal. if an other team member has the goal in the same area with a higher priority, it sends information about its current goal and forces the robot with a lower priority to choose another goal and generate a new plan. description of the software architecture of a single robot performing distributed exploration and communicating with other robots making use of aodv follows in the next paragraphs. while designing the application structure for a single robot it was assumed that the application can operate in two roles: robot and display. the application for a robot role controls a robot as it carries out reading of data from its sensors, produces the robot movement in the terrain, performs path planning and communication with other team members. the application for a display role is used to monitor and display the status of the multi-robot exploration process and can vary from the specific implementation. the application of this type is also a part of the mesh network, but it does not directly participate in the exploration. figure 1. proposed application structure. it is recommended to use aodv routing protocol based on the analysis result of routing protocols 22 vol. 6/2016 framework for ad hoc network communication in multi-robot systems from section 2. the application should have the tools to unicast and broadcast messages for aodv implementation. tcp is the most appropriate protocol for the purposes of sending messages between robots. flooding and sending hello messages can be realised with using less reliable udp, which however supports broadcasting. according to the results of a comparative analysis in section 3 nanomsg recommended for implementation as a mom. thus, the nanomsg library provides the transfer of messages between robots via tcp and uses in-proc for transfer of information between different parts within the application itself. because this library does not support the udp, broadcasting should be implemented by means of language in which the application will be realized. fig. 1 shows the proposed general structure of the application, and describes the relationship between it and the external environment. application operations have different frequency of launching and some of them can be called asynchronously. e.g., plan generation usually runs once per two seconds or on the request from the other robot. a command to control robot movement can be sent every 100 milliseconds. the optimal frequency of sending the heartbeat for the purposes of aodv is 1 second. moreover, it is desirable to receive messages from the other robots instantly. message sending is also necessary to make immediately to reduce the overall time of the terrain exploration. this means that the designed application could not be structured in a single loop. for these reasons, the structure is divided into 5 logical parts, which can be implemented as threads: motion control/visualization, path planning, timer, message processing and inbox messages part. these parts of the application share three data structures: map, current plan and scans archive. map keeps an internal representation of the application’s view about the environment and depends on the selected type of the exploration. in the case of frontier-based exploration, the map will be in the form of the occupancy grid. current plan is the path to a goal selected at this iteration of the exploration process in the form of a polyline. scans archive is a data structure that contains scans received from sensors of any robot of the team. that is, the structure stores local maps created by the robot and local maps created by other robots and transmitted via flooding. the principle of the transfer of local maps between robots during the terrain exploration of a multi-robot team is described in [36]. the application for a robot role is determined by using the motion control thread. it is responsible for receiving the current plan, sending signals to the robot motors for movement to the next selected goal, using the robot sensors to obtain scans, which are then stored in the scans archive and receiving new scans from the scans archive received from other robots. finally, the thread conducts a map update on the basis of these data. if the application is launched for the purpose of visualization of the exploration process, it uses a visualization thread. the application of this type does not create its own scans, so the scans archive contains only local maps received from the robots. this thread uses them to update the map, which is displayed on the screen after that. the path planning thread is the main component of each iteration of the terrain exploration process. this thread selects the next goal for the motion and makes the path to it. this plan is written to the shared structure. the thread uses the map to plan. the application uses a timer to generate the plan after a certain predetermined period. the timer is located in the timer thread. the message processing thread is the main connection between the robot and the rest of the team. it is responsible for sending messages to other robots about a new generated path or a new laser scan. it also receives messages from the inbox messages thread. the inbox messages thread deals with receiving messages from other team members and sending them in the messages processing thread. 5. experimental results for functional testing of the proposed framework, the application of a multi-robot team member was implemented. it is implemented in c++ and simulates the work of an exploration robot for experimental verification of the effectiveness of the proposed application structure. it imitates operations of the real robot equipped with a laser range-finder. the robots are controlled by a simple controller which guides the robot exactly along a planned path with a constant speed. ray casting was used to simulated a laser range-finder with no noise incorporated. behavior and properties of the implemented application were experimentally evaluated and its characteristics were compared with a single-robot exploration. the experiments were conducted using a device which has 32 central processing units (cpu) and has 8 gb of memory. all tests were performed in simulation on the map of an empty terrain that has a size of 20 by 20 meters. each cell of the occupancy grid has a size of 5 by 5 centimeters. it means that the occupancy grid has a size of 400 by 400 cells. various numbers of robots in the team are used to test the loading of the network. the team size is varied from one to ten. the application makes the terrain exploration as a single robot, if the team consists only of one robot. various ranges of communication links are used to test the effectiveness of teamwork. the communication range takes one of the values: 5, 10, 15 or 20 meters. the laser has a range of 1.5 or 2 meters. it allows to verify the functionalities performance with a larger number of steps of the exploration process on the same map. the experiments were performed three times for each system configuration. 23 khilda slyusar, miroslav kulich acta polytechnica ctu proceedings (a). (b). figure 2. dependence of the steps on the team size. the value of the laser range is: (a) 2 meters; (b) 1.5 meters. the time measurement of the exploration is made in steps, i.e the number of calls of the function that changes the robot position. this function is triggered periodically after a certain period of time. thus, a step is a convenient alternative to time. in addition, the usage of steps metric allows to maintain usable statistics of events that pass between sequential steps. 5.1. time of the exploration fig. 2 shows the dependence of the average exploration time for a team member on the number of robots in the team. the graph shows that any team work helps to reduce exploration time. in the beginning, both graphs show a reduction of the number of steps, then there is an increase in steps quantity. this is due to the fact that the robots have to excessively agree on the choice of a goal in bigger teams. as can be seen from fig. 3, a small team disperses at the beginning of the exploration and explores non-overlapping areas of the terrain. however, an excessive number of robots in the team causes the problem of determination of the non-overlapping areas. so robots are required to negotiate the rescheduling of their goals. the graph also shows that the decrease in the communication range increases the number of steps that are required to complete the terrain exploration. some scans from the other robots could not reach the robot because of lack of communication link due to the small value of the communication range. that is, the robot is forced to explore the given area by itself. it increases the overall time of the exploration. comparison of fig. 2(a) and fig. 2(b) shows that the ratio of the steps number and the team size remains constant regardless of the laser range. however, reducing the value of laser range increases the total number of steps. if the laser has a shorter range, then each scan provides less information about the environment. thus, it is necessary to make larger movements (a). (b). (c). (d). (e). (f). figure 3. the resulting paths after the exploration. the laser range of the first column is 2 meters, the second column is 1.5 meters. rows are separated by team size: (a)+(b) 1 robot; (c)+(d) 2 robots; (e)+(f) 5 robots. 24 vol. 6/2016 framework for ad hoc network communication in multi-robot systems and get more scans to obtain a complete map of the terrain. 5.2. plan generations fig. 4 shows the frequency with which the application is launching a new request for a new generation of the plan. a plan generation runs periodically by a timer part of the application, as described in section 4. in addition, it can be run by a request from an another team member. in this implementation, if two robots have the goal at the same area, the robot with a higher priority of the goal sends a request to the robot with a lower priority to select a new navigation goal and generate a plan to the selected goal. the launching frequency of generation of the plan is 20 steps in the application. the graphs shown on fig. 4 contain all plan generations, which had one member of the robot team during complete exploration of the environment. thus, all function values under the value of 20 steps means that the corresponding plan generation is launched on request from an another robot. fig. 4(a) shows that a single robot starts a process of path generation by approximately the same intervals. the increase of the team size leads to the fact that robots can compete for the goals in a particular region of the terrain. this can have both positive and negative effects. let’s consider the situation of the terrain exploration by a team that consists of ten robots and the communication range is 20 meters. this means that all robots agree on the choice of goals with the other robots. the first low part of the graph shown in fig. 4(d) corresponds to fig. 5(a). the robots start the goal selection. however, the current implementation of the goal choice requires to select the goal that does not lie close to the goals of other team members. this means that the distance between the selected goal and goal of another robot should be greater than the laser range. an example of such area is depicted in fig. 5(a) as the yellow circle. if each robot chooses a goal which does not overlap with the goals of the other robots, the plan generation is invoked periodically by the timer thread. that is, the robot does not encounter other robots and it made the right goal choice. the straight part of the plot with a step value, which is equal to 20, corresponds to fig. 5(b). then, most of the terrain is explored and the robots are in a small unexplored part of the environment (see fig. 5(c)). the weak point of the implemented simple prioritizing between the robots appears. at this point, the frequency of plan generation begins to increase. moreover, increasing the number of plan generations leads to increase in the number of steps that are necessary to complete the terrain exploration. it then leads to increase of the total exploration time. 6. conclusion this paper addresses the problem of communication in distributed multi-robot teams forming a mobile ad hoc network. a list of requirements for a routing protocol was drawn up. the comparative analysis of routing protocols has shown that aodv routing protocol is most suitable. in addition, the analysis of three moms was conducted and the nanomsg library was chosen as the most suitable for implementation. multi-robot exploration has been selected as a proofof-concept task. the general structure of the robot application which carries out exploration of an unknown environment in a multi-robot team and which communicates with other robots employing aodv has been proposed. to verify the proposed architecture and offered tools the application has been implemented and has been tested by series of experiments. thus, the paper describes the architecture of the application and possible means of implementation in a generalized form. the specific implementation can be done for different systems and written on a wide range of languages. in future, this application will be used for testing on the group of real robots that solves the problem of the exploration of an unknown environment. goal allocation was not addressed in the paper as it was out of scope of it. analysis and comparison of a variety of sophisticated goal allocation mechanisms will be another stream we would like to go in future. acknowledgements this work has been supported by the technology agency of the czech republic under the project no. te01020197 “centre for applied cybernetics”. access to computing and storage facilities owned by parties and projects contributing to the national grid infrastructure metacentrum, provided under the programme “projects of large infrastructure for research, development, and innovations” (lm2010005), is greatly appreciated. references [1] w. zhigang, l. liu, m. zhou. protocols and applications of ad-hoc robot 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[30] j. broch, d. a. maltz, d. b. johnson, et al. a performance comparison of multi-hop wireless ad hoc network routing protocols. in 4th annual acm/ieee international conference on mobile computing and networking (mobicom), pp. 85–97. association for computing machinery (acm), 1998. doi:10.1145/288235.288256. [31] apache. apollo 1.7.1, 2016. https://activemq.apache.org/apollo/. [32] p. hintjens. zeromq: messaging for many applications. o’reilly media, 1st edn., 2013. [33] m. sustrik. the nanomsg library, 2015. http://nanomsg.org/documentation.html. [34] b. yamauchi. a frontier-based approach for autonomous exploration. in proc. ieee computational intelligence in robotics and automation, pp. 146–151. 1997. doi:10.1109/cira.1997.613851. [35] m. kulich, j. faigl, l. přeučil. on distance utility in the exploration task. in ieee international conference on robotics and automation, pp. 4455–4460. 2011. doi:10.1109/icra.2011.5980221. [36] b. yamauchi. frontier-based exploration using multiple robots. in second international conference on autonomous agents (agents), pp. 47–53. 1998. doi:10.1145/280765.280773. 27 http://dx.doi.org/10.1109/icra.2011.5980367 http://dx.doi.org/10.1109/cicn.2011.129 http://dx.doi.org/10.5120/1270-1749 http://dx.doi.org/10.1109/inmic.2004.1492924 http://dx.doi.org/10.7763/ijiet.2013.v3.223 http://dx.doi.org/10.1145/190314.190336 http://dx.doi.org/10.1145/75246.75269 http://dx.doi.org/10.1007/978-0-585-29603-6_5 http://dx.doi.org/10.1109/infcom.1997.631180 http://dx.doi.org/10.1109/tvt.2013.2279111 http://dx.doi.org/10.1109/icupc.1997.627227 http://dx.doi.org/10.17487/rfc3561 http://dx.doi.org/10.5120/1291-1765 http://dx.doi.org/10.1145/288235.288256 https://activemq.apache.org/apollo/ http://nanomsg.org/documentation.html http://dx.doi.org/10.1109/cira.1997.613851 http://dx.doi.org/10.1109/icra.2011.5980221 http://dx.doi.org/10.1145/280765.280773 acta polytechnica ctu proceedings 6:18–27, 2016 1 introduction 2 ad hoc routing protocols 2.1 proactive routing protocols 2.2 reactive routing protocols 2.3 hybrid routing protocols 2.4 comparison of routing protocols 3 message oriented middleware 4 distributed exploration 5 experimental results 5.1 time of the exploration 5.2 plan generations 6 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0075 acta polytechnica ctu proceedings 13:75–79, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app solid burnt bricks’ tensile strength aneta maroušková∗, jan kubát czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: aneta.marouskova@fsv.cvut.cz abstract. this paper deals with experimental testing of solid burnt bricks and mortar in pure (axial) tension. the obtained working diagrams will be further use for a detailed numerical analysis of whole brick masonry column under concentric compressive load. failure mechanism of compressed brick masonry column is characterized by the appearance and development of vertical tensile cracks in masonry units (bricks) passing in the direction of principal stresses and is accompanied by progressive growth of horizontal deformations. these cracks are caused by contraction and interaction between two materials with different mechanical characteristics (brick and mortar). the aim of this paper is more precisely describe the response of quasi-brittle materials to uniaxial loading in tension (for now only the results from three point bending test are available). for these reasons, bricks and mortar tensile behavior is experimentally tested and the obtained results are discussed. keywords: tensile strength, pure tension, brick, mortar. 1. overview the aim of this paper is determine solid burnt bricks’ and mortar tensile strength in “pure” tension. obtained working diagrams will be further used for detailed numerical analysis of whole masonry brick column. under concentric compressive load of masonry column, bricks are gradually damaged by lateral tension accompanied by appearance and development of cracks. these cracks are caused by contraction and interaction between two materials with different mechanical characteristics as a result of this mutual interaction, mortar that has a lower young’s modulus and has tendency to greater lateral strain, is transversely “pressed” and masonry units, on the contrary are transversely “stretched” [1, 2]. 2. experimental setup for bricks six p20 solid burnt bricks with dimensions of 290×140×65 mm (the declared compressive strength is 20 mpa) were experimentally tested in pure tension test. all specimens were measured and the specific dimension were used for each specimen. steel plates were glued on brick’s face in longitudinal directions by means two components epoxy glue akepox 2040 according to technological procedure in laboratory conditions. akepox 2040 reaches the maximal strength in seven days, after that the specimen were tested. brick with steel plates were placed into testing equipment labtest 4.100sp1 by means two snap hooks (fig. 1). this equipment adds load in form of deformation in regular steps. figure 1. experimental setup. 75 http://dx.doi.org/10.14311/app.2017.13.0075 http://ojs.cvut.cz/ojs/index.php/app aneta maroušková, jan kubát acta polytechnica ctu proceedings figure 2. failure of specimens no. 4 and 5. 3. discussion of the obtained results failure occurred at the interface between steel plate and brick with the exception of two samples. these two samples had initial damage from manufacturing. during experimental testing was observable a slowly opening of the crack following by brittle failure. failure mode of both specimen (no. 4 and 5) is showed in fig. 2. after failure, different level of burning is evident on the cross-section of specimens. grey core in specimen 5 indicates incomplete burning process. specimen 5 showed lower tensile strength, however it is figure 3. cross-sections of specimens no. 4 and 5. figure 4. failure of specimens no. 1, 2, 3 and 6. necessary take into account the size of initial crack from manufacturing. other specimen showed failure between brick and one of steel plates. reason of this type of failure is given by low interface adhesion. usage of insufficient amount of epoxy resin is probably the explanation. another stone specimens were also tested in pure tension test and the expected failure was reached – all specimens failed in stone (interface strength was higher than specimen tensile strength). fig. 4 shows failure of specimens 1, 2, 3 and 6. 76 vol. 13/2017 solid burnt bricks’ tensile strength figure 5. working diagram force-elongation of bricks. figure 6. tensile strength of bricks. figure 5 shows working diagrams of each specimen. the highest value of force was reached for specimen 6. specimen 6 (together with specimen 3) was quite enough glued (see fig. 4) and the failure occurred predominantly in very first layer of masonry unit [3]. this layer has improved properties thanks to glue penetration into porous system. for this reason, there is no failure direct at the interface, but a small layer of brick was thrown away. specimen 1 and 2 show poor gluing of steel plate, especially specimen 2 has the lowest value of reached force. specimens 4 and 5 had an initial manufacturing crack and have failed in brick. the maximal value of reached force is different, it could be caused by distinct depth of initial crack. it was not possible to measure the size of initial cracks, because there were only hair cracks observed. the possible solution is using specimen with notches whose accurate size would enable a conversion force to area. the initial slip of the specimens is not displayed in working diagrams (fig. 5). the diagrams show brittle rupture in all cases. average tensile strength was determined to figure 7. young’s modulus of bricks. figure 8. failure of mortar specimens. 0.64 mpa with standard deviation 0.15 mpa. specimen no. 2 was not taken into account. young modulus was obtained as secant modulus between two loads: 90 % num, 40 % num, where num is maximal load. average modulus of elasticity was determined to 0.24 gpa with standard deviation of 0.03 gpa. specimen no. 2 was not taken into account. 4. experimental setup for mortar for mortar pure tension test, three mortar specimens with dimensions 40×40×160 mm were made. specimens were tested after 28 days to ensure mortar’s hardness. steel plates were glued on mortar specimens after 21 day from construction (meaning the maximal strength of epoxy resin in 7 days was respected). mortar with steel plates were placed into testing equipment labtest 4.100sp1 by means two snap hooks as well as bricks. 77 aneta maroušková, jan kubát acta polytechnica ctu proceedings figure 9. failure of specimen m2. figure 10. working diagram force-elongation of mortar. 5. discussion of the obtained results from mortar testing two mortar specimens show failure close to the glued steel plates. after failure of specimen no. m1 small place of incomplete gluing is evident. specimen no. m2 failed approx. in a half of its length. calculated values of tensile strength and young modulus are reported in bar charts (fig. 11 and 12). the initial slip of the specimens is not displayed in working diagrams (fig. 10). the diagrams show brittle rupture in all cases. average tensile strength was determined to 0.25 mpa with standard deviation 0.04 mpa. young modulus was obtained as secant modulus between two loads: 90% num, 40% num, where num is maximal load. average modulus of elasticity was determined to 0.047 gpa with standard deviation of 0.005 gpa. figure 11. tensile strength of mortar. figure 12. young’s modulus of mortar. 6. summary reached values of young’s modulus and tensile strength of bricks are lower than the expected. the expected tensile of bricks strength was approx. around 1.5 mpa and the reached value with the exception of specimen 6 is not even half of that. simultaneously stone specimens were tested, whose failure matches to expected, and therefore the amount of used epoxy resin was determined as the reason of lower obtained values. also the low tensile strength of the specimens with the manufacturing crack (no. 4 and 5) for which were not possible to determine the effective area resisting to the force and thus it was not possible to recalculate the obtained tensile strength for the effective area. after this recalculation, the value of tensile strength of specimens 4 and 5 is expected higher than the determined ones. the experimental testing of bricks will be repeated and the results mutually compared and discussed. tensile strength of mortar met the expectation; however due to small amount of specimens the testing will further continue and reached values will be evaluated. acknowledgements the results were obtained with support from research project df12p01ovv037 – “progressive non-invasive methods of the stabilization, conservation and strengthening of historic structures and their parts with fibreand nanofibre-based composite materials”. the grant researcher is prof. ing. jiří witzany, drsc. 78 vol. 13/2017 solid burnt bricks’ tensile strength references [1] j. witzany, t. čejka, t. zigler. failure mechanism of compressed short brick masonry columns confined with frp strips. construction and buildings materials 63:180– 188, 2014. doi:10.1016/j.conbuildmat.2014.04.041. [2] a. maroušková. masonry column reinforced by frp wrapping: behavior and numerical analysis. applied mechanics and materials 825:27–30, 2008. doi:10.4028/www.scientific.net/amm.825.27. [3] r. fedele, g. milani. a numerical insight into the response of masonry reinforced by frp strips. composite structures 92:2345–2357, 2010. doi:10.1016/j.compstruct.2010.03.014. 79 http://dx.doi.org/10.1016/j.conbuildmat.2014.04.041 http://dx.doi.org/10.4028/www.scientific.net/amm.825.27 http://dx.doi.org/10.1016/j.compstruct.2010.03.014 acta polytechnica ctu proceedings 13:76–80, 2017 1 overview 2 experimental setup for bricks 3 discussion of the obtained results 4 experimental setup for mortar 5 discussion of the obtained results from mortar testing 6 summary acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0035 acta polytechnica ctu proceedings 13:35–38, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app foamed cement composites: detection of the modulus of elasticity using dic analysis and comparison with other methods jakub ďurejea, ∗, zdeněk prošeka, b a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b university centre for energy efficient buildings, czech technical university in prague, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: jakub.dureje@fsv.cvut.cz abstract. a modulus of elasticity was determined for eight differently foamed cement paste samples. samples were loaded in the laboratory by a hydraulic press. the force acting on the sample was read directly from the laboratory press. digital image correlation (dic) analysis were used to draw deformations. before loading pressure test was applied a random contrast pattern to the samples. samples were captured by the camera in a one-second interval during the loading pressure test. the images were edited in the adobe photoshop lightroom and then evaluated using ncorr software. the result is a vertical and horizontal shift field. on the basis of the results obtained, it was possible to calculate the modulus of elasticity of each sample. keywords: dic analysis, foamed cement paste, field of displacement, static modulus of elasticity, dynamic modulus of elasticity. 1. introduction there are many different methods to detect deformation of the sample during the loading pressure test. the displacement of the crosspiece of the hydraulic press does not correspond to the deformation of the sample in the direction of the force. this is due to the deformation of the exponed parts of the test device. to measure deformations, it is possible to use some type of extensometer and place it directly between the jaws of the press in which the sample is tested and measure the displacement directly between the jaws. thanks to the ever-improving technologies, we can obtain the entire field of displacement by dic analysis by capturing a series of photos during sample testing and subsequent evaluation. the advantage is that, compared to the measurement with an extensometer, we obtain deformations in all directions on the measured surface of the sample [1, 2]. dic analysis is used in many industries, by many companies, including airbus, awe, stresscraft and british energy. it is used, for example, for measuring the heat dissipation and curvature of electronic components, for measuring the 3d shape of the air bags, for monitoring the components of the power plants to determine their status and to reduce their emissions, or for the behavior of building materials during the loading tests [2]. dic analysis is an optical method, to get the most accurate results it is necessary to avoid optical defects. for each lens, the image is distorted towards the edges. depending on the deformation of the square, we distinguish distortions to barrel distortions, pincushion distortions, or a combination of both – wave distortions. lens manufacturers are trying to eliminate these phenomena with the appropriate lens system. the lenses are constructed at certain distances where the distortion occurs least. for this reason, it is necessary to select a suitable lens and capture distance according to the size of the subject to be sensed [3, 4]. 2. materials and samples the sample properties were measured at 28 days of age and the day after manufacture the samples were removed from formwork and stored freely in a laboratory environment at a temperature of 22 ± 1 ◦c and a relative humidity of 50 ± 2%. test beams of dimensions 40 × 40 × 160 mm were made, which were broken in half during the previous bend test. the reference mixtures for testing were made of portland cement cem i 42.5r radotin (refc), a combination of portland cement and finely ground concrete recyclate in a ratio of 0.66 : 0.34 (refr). fine-grained recycled concrete was produced by high-speed grinding, from waste generated from drainage concrete blocks. subsequently, the reference mixtures were lightened using foaming additives such as industrial additive from the mapei (m) company or the additive used by firefighters (h). the composition of the individual samples and the bulk density are given in table 1 [5]. 3. experimental methods dic analysis was performed on a total of 10 samples, on two references and eight differently foamed cement pastes. samples were loaded in the laboratory by a 35 http://dx.doi.org/10.14311/app.2017.13.0035 http://ojs.cvut.cz/ojs/index.php/app jakub ďureje, zdeněk prošek acta polytechnica ctu proceedings sample nr. designation cement (c) recyclate (r) weight (c+r) water to mix cement putty (v1) water added to the foam (v2) water (v1+v2) w/c foam bulk density [g] [g] [g] [g] [ml] [ml] [-] [ml] [kg.m−3] 45 c7.5h 750.0 0.0 750 225 75 300 0.40 3.75 603.83 40 c7.5m 750.0 0.0 750 225 75 300 0.40 3.75 677.84 29 c10h 1000.0 0.0 1000 300 50 350 0.35 2.50 1333.10 21 c10m 1000.0 0.0 1000 300 50 350 0.35 2.50 1394.40 24 r7.5h 487.5 262.5 750 225 75 300 0.40 3.75 1177.08 34 r7.5m 487.5 262.5 750 225 75 300 0.40 3.75 895.14 15 r10h 650.0 350.0 1000 300 50 350 0.35 2.50 1603.99 12 r10m 650.0 350.0 1000 300 50 350 0.35 2.50 1457.30 5 refc 1000.0 0.0 1000 350 0 350 0.35 0.00 1646.10 9 refr 650.0 350.0 1000 385 0 385 0.39 0.00 1641.70 table 1. composition of the samples and the bulk density [5]. figure 1. random pattern on a sample in grayscale. hydraulic press heckert model fp100, at a speed of 1 mm per minute. subsequently, a static modulus of elasticity was specified. after that, static modulus of elasticity obtained on the basis of dic analysis, static modulus of elasticity obtained on the basis of the displacement of the hydraulic press and the dynamic elastic modulus obtained by the ultrasonic and resonance method were compared. nondestructive resonance method was used to measure dynamic modules. the brüel & kjær measuring set consists of the figure 2. field of vertical displacement from the ncorr post software. brüel & kjær type 3560-b-120 measuring switchboard, the brüel & kjær type 4519-003 acceleration sensor, the brüel & kjær type 8206 impact hammer and the control laptop. modulus of elasticity obtained from data from crosspiece of the hydraulic press does not correspond to the reality, because the deformation of the test instrument itself is included in the data. also, the values of the dynamic modulus of elasticity are rather inaccurate due to the inappropriateness of the resonant and ultrasonic method on porous systems. dynamic modulus of elasticity should achieve higher values than static modulus of elasticity, because by non-destructive methods we get the tangent modulus 36 vol. 13/2017 foamed cement composites: detection of .... figure 3. results of static and dynamic moduli of elasticity. of elasticity and the secant modulus of elasticity we get from the loading test [6]. 4. princip of dic analysis dic (digital image correlation) is an optical method used to measure displacement on the sample surface. compares the grayscale values of individual groups of points called subsets on the surface of the sample between the pictures taken during the loading test. it is necessary to have a sufficiently contrasting random pattern on the surface of the sample. some materials have this pattern naturally. for other samples, the pattern should be applied appropriately. the software compares shades of gray and their displacement between individual photos (figure 1). the result is the field of displacement between the reference and each picture. the reference picture is usually taken before the start of the test [7, 8]. 5. advantages of dic analysis dic analysis is a cost-effective optic method that can be used in laboratory and outdoor environment. it is used to inspect large important structures (bridges, industrial buildings, power plants) that degrade due to external influences. the capture of large objects can be done directly from the air using drones thanks to advanced technologies. this solution is relatively simple, accurate and financially advantageous [1, 2]. during the measurements in the laboratory, we obtain a field of displacement in the horizontal and vertical directions. we can also measure cracks occurring. during the loading tests, the exact course of cracks can be determined by dic analysis, even those that are not observable by the naked eye. these data are very beneficial for the presumption of the durability of the tested material [2]. 6. sample measurement dic analysis was used on foamed cement paste samples. half of cement paste beams were tested according to csn en 12390-3 test of hardened concrete part 3: compressive strength of test samples [9]. a random contrast pattern of black and white spray was applied to the samples. during the testing, images were taken with canon eos 70d on a tripod from a distance of approximately 300 millimeters. the images were taken in a one-second sequence during loading. for each sample, 50 − 100 photos were taken. the photos were edited in adobe lightroom for higher contrast, and then the dic analysis was performed in ncorr version 1.2.1. the results were scaled and the virtual extensometer was placed in the field of displacement in the ncor_post software (figure 2). the results of the individual displacement were assigned to the respective forces using the extractforces software. subsequently, a static modulus of elasticity in microsoft excel was calculated from the results. 7. results and discussions samples ref c and ref r were references. the other eight samples were differently foamed. samples 37 jakub ďureje, zdeněk prošek acta polytechnica ctu proceedings c 7.5 h, c 7.5 m, and r 7.5 m were foamed too much therefore it was not possible to measure the dynamic modulus of elasticity by the resonance method because it was not possible to capture the response from the impact hammer. the results of the other modulus of elasticity in these three samples, because of their high foaming, are not very accurate. in the case of the compression test, these three samples were observed (especially for samples c 7.5 h and c 7.5 m), as the material collapsed successively on the floors. it was not about the flexible behavior that is assumed to acquire the modulus of elasticity. therefore, the results of the measurements of these samples should be considered inaccurate, and due to their poor mechanical properties, do not use this highly foaming cement paste for practical use in the building industry. the static modulus of elasticity for all samples obtained from the displacement of the hydraulic press crosspiece are very low. this is due to the deformation of the test rig itself together with the sample deformation. for this reason, it makes no sense to further consider the static modulus of elasticity thus obtained. measurement of c 10 m, r 10 h and r 10 m samples were expected, static moduli of elasticity are slightly less than dynamic. for samples r 7.5 h, ref c, and ref r, the static moduli of elasticity are approximately half that of the dynamic moduli of elasticity. on the other hand, the static modulus of elasticity is higher than the dynamic for c 10 h sample (figure 3). this is due to the inaccuracy of measuring the dynamic modulus of elasticity because of the inappropriateness of the resonant and ultrasonic method on porous systems. 8. conclusion for this type of samples, the most accurate method is dic analysis. measurement by ultrasonic method or resonance method is not accurate in porous systems. measurement of deformation directly from the hydraulic device is very inaccurate as it involves deformation of the hydraulic device itself. the results show that the less or no foamed samples without recyclate have a higher modulus of elasticity than samples with recyclate, but with more foamed samples, it is the opposite. furthermore, it is noticeable that the less foamed samples with a foaming additive used by firefighters have a higher modulus of elasticity than samples with a foaming additive for industrial use, but with more foamed samples, this is the opposite. acknowledgements this outcome was supported by the czech technical university in prague under no. sgs16/201/ohk1/3t/11 and no. sgs17/060/ohk1/1t/11 and by the ministry of education, youth and sports within national sustainability programme i, project no. lo1605. the authors also thank the center for nanotechnology in civil engineering at the faculty of civil engineering, czech technical university in prague, and the joint laboratory of polymer nanofiber technologies of the institute of physics, academy of science of czech republic, and the faculty of civil engineering, czech technical university in prague. references [1] p. k. rastogi, e. hack. optical methods for solid mechanics: a full-field approach. john wiley & sons, 2012. [2] n. mccormick, j. lord. digital image correlation. materials today 13(12):52–54, 2010. [3] jasar. optické vady a vlastnosti objektivu. http: //photo.mysteria.cz/clanky/objekt2.html#aberace, 2002. [4] l. jelinek. darktable: korekce objektivu a odstranění skvrn. https://www.linuxexpres.cz/praxe/ darktable-korekce-objektivu-a-odstraneni-skvrn, 2015. [5] v. růžička. mechanické vlastnosti pěnových cementových pas. master’s thesis, czech technical university in prague, prague, 2017. [6] m. ballon. porovnaní modulů pružnosti stavebních materialů. master’s thesis, brno university of technology. faculty of civil engineering, brno, 2014. [7] m. a. sutton, j. j. orteu, h. schreier. image correlation for shape, motion and deformation measurements: basic concepts, theory and applications. springer science & business media, 2009. [8] s. yoneyama, g. murasawa. digital image correlation. experimental mechanics 2009. [9] čsn en 12390-3–zkoušení ztvrdlého betonu–část 3: pevnost v tlaku zkušebních těles. úřad pro technickou normalizaci, metrologii a zkušebnictví 10, 2009. 38 http://photo.mysteria.cz/clanky/objekt2.html#aberace http://photo.mysteria.cz/clanky/objekt2.html#aberace https://www.linuxexpres.cz/praxe/darktable-korekce-objektivu-a-odstraneni-skvrn https://www.linuxexpres.cz/praxe/darktable-korekce-objektivu-a-odstraneni-skvrn acta polytechnica ctu proceedings 13:36–39, 2017 1 introduction 2 materials and samples 3 experimental methods 4 princip of dic analysis 5 advantages of dic analysis 6 sample measurement 7 results and discussions 8 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0097 acta polytechnica ctu proceedings 13:97–101, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app geometrical optimization of dental implants with regard to osseointegration luboš řehounek∗, františek denk, aleš jíra czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: lubos.rehounek@fsv.cvut.cz abstract. a newly developed reference dental implant specimen type was subjected to numerical simulations of osseointegration. the goal of these tests was to optimize the geometry of the implant so as to reduce local stress concentrations and provide a better flow of stress through the whole implant body. conditions for osseointegration were considered when evaluating the anchoring system of the implant in regard to its placement in the human cancellous and cortical bone. numerical simulations showed that stress concentrations occur mostly in the upper cylindrical part of the implant. by increasing the width of this cylindrical part, we were able to reduce the maximum values of von-mises stress by 20 %. keywords: dental implant, optimization, trabecular, numerical, fem. 1. introduction the main shortcoming of commonly manufactured dental implants is their predisposition towards early failure [1–3]. this unwanted factor can be attributed to uneven distribution of stresses in the implant, periimplant area and bone [4, 5]. our main goal in developing new implants was to reduce stress concentrations occuring mainly in the upper cylindrical part of the implant, just below the abutment, while preserving the osseointegration benefits of the implant. the tested implant has been patented in the czech republic [6]. the numerical analysis describes the mechanical behavior of the implant in regard to mechanical loads and the corresponding biomechanical stablity in the bone tissue. the tests focus on two aspects – mechanical behavior of the implant under a prescribed mechanical load and osseointegration of the implant inside the human mandible. the first part of the analysis focuses on the osseointegration of the implant in human mandible. the aim of the second part of the tests was to determine the ultimate bearing capacity of the dental implant stem according to the čsn en iso 14081 – “stomatology – implants – dynamic fatigue testing of intraosseous dental implants” standard. this paper will only focus on the first part of the tests (osseointegration) and leave the second part (mechanical tests) for further studies, as some results are already available, but additional simulations are still needed. the analysis provides an overview of stress and deformation relations of the tested implant stem in a prescribed environment in all of its parts. this is simulated by embedding the implant stem in a block of human bone (fig. 2) and subsequently evaluating the stress and deformation relations at surface areas of the stem and at the interfaces of the stem and bone tissue. numerical simulations were done in the ansys apdl system by incorporating previously created 3d geometries (fig. 1) into the environment of the software. the geometry of the implant stem was created in detail at the surface areas of the implant. the internal part of the stem has been created with less detail (i.e.: without the screw-thread). 2. numerical testing of osseointegration for the purpose on numerical analysis in ansys apdl, we created an environment where the dental implant is fixed into the bone tissue of human mandible. the implant consists of the intraosseous stem and the implant superstructure, which represents the abutment and the crown. the connection of individual volumetric bodies is attained using rigid constraints generated by the volumetric “vglue” operations. the mesh of the model consists of quadratic tetrahedrons type solid 187 generated in the “smart size” mode. the material of all components of the dental implant is the titanium alloy ti-6al-4v. the stem of the implant is fixed into a volumetric model of human mandible, situated according to real implantation procedures. the geometrical orientation of the ribs was arranged in the least favorable position (in regard to the results of mechanical tests). the exact geometry of the bone was generated by boolean operations (subtracting the full body volume of the stem from the body of the bone). the connection between contact surfaces of the stem and the bone was attained by volumetric “vglue” operations. the numerical model of the mandible was then subsequently remeshed according to the new boundary conditions (fig. 2). the front faces of the mandibles have symmetric 97 http://dx.doi.org/10.14311/app.2017.13.0097 http://ojs.cvut.cz/ojs/index.php/app l. řehounek, f. denk, a. jíra acta polytechnica ctu proceedings figure 1. 3d model of the designed implant 0001 (ansys apdl): whole 3d model (left), longitudinal section (middle) and fem mesh (right). figure 2. numerical model of the mandible block which supports the intraosseous part of the implant stem – volumetric model (left), fem mesh (middle) and total deformation isolines (right). boundary conditions and the lower strip of nodes is set to have zero displacement value. physiological load is represented by a single slant force applied to the central point of the upper base of the superstructure. the force magnitude is f = 130 n. the direction of the force is defined by contribution of individual force components according to the total load imposed on a single tooth. the magnitude of normal force is 125 n and tangent force is 35 n (fx = 25 n, fy = 25 n, fz = 125 n). 3. optimization of the cylindrical part the following optimization will assume that the small strip of the cylindrical stem forming the upper structure of the stem is always the area where values of stress concentrate. these extremities are caused figure 3. numerical analysis of the reference implant 0001: displacement – umax = 25 µm (left), equivalent von mises stress – σeqv,max = 91.7 mpa (middle and right). 98 vol. 13/2017 geometrical optimization of dental implants figure 4. optimization of the geometrical shape in the upper cylindrical part: reference model 0001a (cyl. width 0.50 mm) (left), model 0001b (cyl. width 1.25 mm) (middle) and model 0001c (cyl. width 3.00 mm) (right). figure 5. numerical analysis of the reference implant 0001b: displacement – umax = 25 µm (left), equivalent von mises stress – σeqv,max,1 = 97.1 mpa, σeqv,max,1 = 73.2 mpa (middle and right). by complex geometrical shapes and also a relatively small width of the cylindrical strip. the goal of the optimization will be to effectively modify the geometrical shape of the implant in the area of fixation of the stem into cortical bone of the human mandible. also, the overall stability of the intraosseous part must not be compromised. two new variants have been tested – variant 0001b and 0001c. the modification represents increasing of the width of the cylindrical part. the width of the cylindrical part of the reference implant 0001a is 0.50 mm. the model 0001b had the width increased up to 1.25 mm and model 0001c up to 3.00 mm. identical simulations were run on all of the mentioned variants and the outcomes of the numerical analyses are shown in fig. 3, fig. 5 and fig. 6. on the basis of equivalent von-mises stress disfigure 6. numerical analysis of the reference implant 0001c: displacement – umax = 25 µm (left), equivalent von mises stress σeqv,max = 73.6 mpa (middle and right). tribution at the surface areas of the imlpant stem, we can observe that the widening of the cylindrical strip of the upper part without any other geometrical modification positively affects the stress distribution by lowering extremities. the variant 0001b (cylindrical part 1.25 mm) has a mean value of σeqv,avg = 61.6 mpa and the stress flow is distributed over a larger area. there is an extreme of σeqv,max = 73.2 mpa. there is also another extreme of σeqv,max = 97.1 mpa spread over a very small area situated in the edge of the cylindrical part. the system of vertical ribs is situated in both cortical and cancellous bone. the variant 0001c (cylindrical part 3.00 mm) has a maximum value of σeqv,max = 73.6 mpa, but the system of vertical ribs is situated now only in cancellous bone. 99 l. řehounek, f. denk, a. jíra acta polytechnica ctu proceedings implant type osseointegrated in bone tissue of mandible displacement of the implant and bone tissue equivalent von-mises stress optimization influence ut [µm] σeqv,max [mpa] [%] implant 0001a 25 91.70 100.00 implant 0001b 25 73.20 79.80 implant 0001c 25 73.60 80.00 table 1. overview of the results provided by geometrical modification. reference implant 0001a and modified implants 0001b and 0001c. figure 7. two stl files showing the final geometrical solution of the implant (trabecular outer shell second from left) (left), the final implant model (middle) and homogenized layer of the outer trabecular structure for easier simulations (right). 4. conclusions 4.1. results the results of the numerical simulations show that by increasing the upper cylindrical portion of the stem, we can improve the distribution of stresses at the surface areas of the implant and also reduce extreme values of stress. the variant 0001b showed that by increasing the width of the cylindrical part by 0.75 mm (from 0.50 mm to 1.25 mm), we can reduce stresses by approximately 20 % and achieve a more even distribution. there is, however, an extreme value of 97.1 mpa situated at the edge of the cylindrical part over a very small area. this extreme is caused by discontinuities in geometry (sudden change of curvature), which will be accounted for by rounding the edge of the cylinder by a radius of at least 0.1 mm during the production process. biomechanical stability is achieved by situating the vertical ribs of the implant inside the cortical bone of the mandible. the variant 0001c (width from 0.75 mm to 3.00 mm, 2.50 mm increase) also showed a 20 % reduction in values of stress. the cylindrical part of the implant does not permeate through the whole layer of cortical bone, so the distribution of stress is more even and there are no larger extremes in values of stress. this geometrical solution, however, might not provide sufficient biomechanical stability as the vertical ribs, which are supposed to anchor the implant in the bone, are only situated in cancellous bone. considering the biomechanical implications of the presented results, we evaluated the variant 0001b with a 1.25 mm wide cylindrical strip as the most efficient solution, which reduces extreme values of stress as well as provides sufficient stability inside thevbone tissue. 4.2. future prospects the chosen variant 0001b has good mechanical stability and stress distribution and was chosen as the variant for further development. with the shape solution of the implant stem complete, the next step in the research will be to create a numerical model of a composite implant (fig. 8). this implant will comprise of both homogeneous and trabecular portions. trabecular structure has been discussed at length in one of other author’s works [7]. the implant will maintain its homogeneous stem. the outer shell with the stabilizing ribs will, however, be comprised exclusively of the trabecular structure. further simulations will be made and multiple variants of different layer thickness analysed in order to determine which shape solution is the best for osseointegration, mechanical stability and machanical bearing capacity. the approach towards modelling the trabecular structure in future research will be focused on sustainable, repeatable analyses. we have already found this 100 vol. 13/2017 geometrical optimization of dental implants figure 8. whole implant body of the implant for further research. implant has a homogeneous stem and an outer layer comprising of the trabecular structure. the implant has been patented in czech republic [6]. to only be possible by making a virtual homogenized material (fig. 7) with specific curve-fitted properties determined by experiment [7]. this approach allows us to apply previously determined mechanical properties of the trabecular structure onto a homogeneous material that is easy to analyze in numerical environment. by running a few simple analyses, we found this to be the optimal approach as it is not very demanding in computation times and it maintains sufficient precision. by manipulating the shape of the outer trabecular structure, we will then be able to determine which variant is optimal in regard to reduction of stress shielding and mechanical bearing capacity. acknowledgements the financial support by the faculty of civil engineering, czech technical university in prague (sgs project no. sgs17/168/ohk1/3t/11) is gratefully acknowledged. references [1] l. le guéhennec, a. soueidan, p. layrolle, y. amouriq. surface treatments of titanium dental implants for rapid osseointegration. dental materials 23(7):844–854, 2007. doi:10.1016/j.dental.2006.06.025. [2] r. osman, m. swain. a critical review of dental implant materials with an emphasis on titanium versus zirconia. materials 8(3):932–958, 2015. doi:10.3390/ma8030932. [3] t. jemt, m. olsson, f. renouard, et al. early implant failures related to individual surgeons: an analysis covering 11,074 operations performed during 28 years. clinical implant dentistry and related research 18(5):861–872, 2016. doi:10.1111/cid.12379. [4] m. i. z. ridzwan, s. shuib, a. y. hassan, et al. problem of stress shielding and improvement to the hip implant designs: a review, 2007. doi:10.3923/jms.2007.460.467. [5] m. long, h. j. rack. titanium alloys in total joint replacement–a materials science perspective. biomaterials 19(18):1621–1639, 1998. doi:10.1016/s0142-9612(97)00146-4. [6] f. denk, f. denk, a. jíra. a dental implant shaft, czech republic, 306457, 2017. [7] l. řehounek, a. jíra. experimental and numerical analyses of a 3d-printed titanium trabecular dental implant. acta polytechnica 57(3):218–228, 2017. doi:10.14311/ap.2017.57.0218. 101 http://dx.doi.org/10.1016/j.dental.2006.06.025 http://dx.doi.org/10.3390/ma8030932 http://dx.doi.org/10.1111/cid.12379 http://dx.doi.org/10.3923/jms.2007.460.467 http://dx.doi.org/10.1016/s0142-9612(97)00146-4 http://dx.doi.org/10.14311/ap.2017.57.0218 acta polytechnica ctu proceedings 13:98–102, 2017 1 introduction 2 numerical testing of osseointegration 3 optimization of the cylindrical part 4 conclusions 4.1 results 4.2 future prospects acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.16.0037 acta polytechnica ctu proceedings 16:37–39, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app soil reinforcement and ec7 (202x) martin vaníček geosyntetika ltd., nikoly tesly 3, 160 00 prague, czech republic correspondence: mvanicek@geosyntetika.cz abstract. soil reinforcement is not really covered in current version of eurocode 7 and therefore during the proposals for 2nd generation of ec7 the soil reinforcement was highly promoted for inclusion in new ec7 based on the comments from practising engineers. in new ec7 which is now being written in newly created part 3 there will be next to other clauses also complete clause on reinforced ground. the paper will inform on the proposed contents of this clause. keywords: soil reinforcement, eurocode 7. 1. introduction existing eurocode 7 covers soil reinforcement topic very vaguely. together with ground improvement topic it is written just on half page [1, 2]. once current eurocode 7 was introduced to praxis, this state was highly criticised by geotechnical engineers around the whole europe. this criticism was formally visible as outcome from periodic 5-year review. as a result of whole eurocode suite success, european commission decided already in 2010 to update and further evolve the eurocode suite towards its second generation. 2. development in between 2010 and 2015 in soil reinforcement area within the cen technical committee tc250/sc7, which is responsible for eurocode 7, fourteen evolution groups (eg) were founded. each eg concentrated on different specific subject of current ec7. eg 5 under the convenorship of the author focused on the introduction of the subject of reinforced soils into the eurocode 7. group members were from all around europe. within this period comprehensive background document was developed, which covers most of typical applications of reinforced soil. we concentrated on specification of relevant limit states that should be checked during the design of such structures. calculation methods for different limit states were discussed and compatibility of these methods was checked. the main differences were identified in the design approaches that were selected in specific countries. surprising recognition was combination of different design approaches used during checking of identified limit states even within one country. in some countries the former codes of praxis were combined with eurocode 7, even when it meant combination of limit states approach with overall stability approach and hence very high (pointless) safety margins were required. during this phase group was meeting face to face once a year within ec7 workgroup meetings, which were held around europe (e.g. in november 2014 in prague). in the meantime, the group held online meetings in about 7 times a year, where further advancement and tasks for periods in between meetings were discussed. it was agreed to focus in turn to the following types of reinforced soils: • reinforced embankment slopes • soil nailed slopes • walls from reinforced soil • subsoil reinforcement under embankments on soft soils • embankment on load transfer platforms over inclusions • embankments over possible voids • veneer stability the structures that were agreed not to cover at the moments are reinforcement under shallow foundations and working platforms reinforcement. within ultimate limit states following locations of limit states origin were identified: • in soil • in reinforcing element • on the interface between soil and reinforcing element • at facing element • the structure as a whole it is possible to illustrate the identified limit states for given structure in the following figure 1. for these limit states it will be necessary to define partial safety factors on both materials and resistances according to the design cases selected in particular country. 37 http://dx.doi.org/10.14311/app.2018.16.0037 http://ojs.cvut.cz/ojs/index.php/app martin vaníček acta polytechnica ctu proceedings retaining structures embankment base piled embankment voids overbridging veneer stability uls in soil stability along slip surface (external, internal and compound) x x x x sliding x x x bearing capacity (including the excentricity) x squeezing x x in reinforcement tensile rupture geosynthetic reinforcement x x x x x steel reinforcement x x1 x1 x1 x soil nails x anchors x2 shear rupture soil nails x rigid vertical bearing elements “rvbe” (piles, micropiles, …) x2 x2 tensile continuity (overlaps, connections, …) x x x x x on the interface between soil and reinforcement pull-out x x x x x stripping x x x sliding x x x x squeezing soil nails x rigid vertical bearing elements “rvbe” (piles, micropiles, …) x2 x2 on facing elements connection to reinforcing element x bulging (blocks, steel mesh, wrap around, …) x sls in reinforced structure itself total vertical / horizontal deformation x x x x x relative vertical / horizontal deformation post construction creep x x x x x relative deformation of the facing (bulging) x in reinforcing element total strain (elongation) x x x x x creep strain (post construction) x x x x x in facing element differential settlement along the facing due to subsoil deformation x notes: x1 rarely used x2 anchors + rvbes are not considered as soil reinforcements in current standards but are shown as an aide memoire for hybrid structures figure 1. matrix of identified limit states and reinforced soil structures. with respect to the need for definition of design strength of reinforcing element, there will be in the framework of second generation of eurocode 7 defined a procedure for determination of strength for geosynthetic reinforcing elements, steel reinforcing elements and soil nails. the evolution group suggested following procedures to be incorporated into the ec7. an agreement between sc3 and sc7 was made that tensile strength of steel soil reinforcing elements will be covered in 2nd generation of ec7 rather than in ec3 [3]. for steel reinforcing elements uk and french long-term experience and suggestions will be used. for geosynthetic elements the design strength will be adopted from iso/tr 20432:2007 [4] that will be further extended to incorporate additional reduction factors from german specification ebgeo [5]. for soil nails design strength current appendix to en 14490 (execution of soil nailed structures) will be adopted [6]. 3. eurocode 7 development since 2015 the actual development of eurocode 7 2nd generation started in 2015 when project teams 1 and 2 were established. team 1 was contracted to reorganize eurocode 7 into 3 parts with common structure in all main clauses which will mimic the structure of other structural eurocodes. part 1 will be general and will be common to all geotechnical structures, part 2 will cover ground investigation and part 3 will cover all different geotechnical structures. additional target for team 1 was to propose simplifications to current text. all team 1 tasks are finished as of september 2017. task of team 2 was to rewrite part 1 of eurocode 7, so it includes the general subjects common to most of geotechnical structures and to prepare an alternative to existing design approaches, which generate incompatibility in between individual countries with respect to design. at the same time the requirement was to make the eurocode 7 fully compatible with eurocodes 0 and 1, which specify the basis for limit states design and loadings on structures. at the end 38 vol. 16/2018 soil reinforcement and ec7 (202x) of october 2017 team 2 prepared the final draft of ec7 part 1 for final comments by cen national standards bodies. these comments have to be taken on board for the final part 1 deliverable at the end of april 2018. second and third part of eurocode 7 started to be updated in october 2017 as during the 2017 the project teams 3, 4 and 5 were selected and appointed. team 3 is responsible for updating of ec7 part 2 for ground investigations. this part will be redrafted in that way that the main focus will be on determination of soil properties needed for geotechnical design from different investigation techniques, rather than the current situation when the focus is on tests. as part of team 3 work will be specification of ground investigation requirements and outputs including reporting. team 4 is responsible for only selected clauses of part 3, which cover the subjects of slopes, shallow and deep foundations and ground improvement. these clauses white spaces will be filled in, current text will be amended, refined and simplified to achieve the ease of use target. the clause on ground improvement will be completely new. team 5, of which the author is member, is responsible for remaining clauses on geotechnical structures in ec7 part 2. more specifically these clauses are retaining structures, anchors and reinforced ground. the task of this team is also to fill in white spaces, amend, refine and simplify existing text, so it will be in accordance with basic structure of all eurocodes. this common structure of all eurocodes should lead to simpler navigation through all eurocodes parts. the plan is that teams 3 to 5 shall have ready in april 2019 first proper draft of their respective clauses. this draft will be than reviewed by tc250/sc7 committee and official comments submitted by them to project teams. after that project teams will prepare final draft by the end of october 2019. the october 2019 draft will be circulated to national standards bodies for official comments from cen member states. these comments will lead to final text of respected parts by project teams by the end of april 2020, when their tasks end. there is also planned 3rd phase of ec7 updating. this phase will be undertaken by project team 6, whose task is to amend the text of all 3 parts for rock mechanics and dynamic design. the schedule for project team 6 is 1 year behind project teams 3, 4 and 5, which means that they should conclude their work by april 2021. once all project teams conclude their work the ownership of eurocode 7 text will be transferred to tc250/sc7 committee, which will undergo the official process of acceptance voting by cen processes. subsequently the accepted text not only of eurocode 7 but of all eurocodes will be published by cen, which is not expected before the end of 2022. the clause on reinforced geotechnical structures will be based on output from evolution group as described above. the main task will be to define the main limit states for different types of reinforced soil structures as indicated in figure 1 and within informative annexes there will be suggested calculation methods for evaluation of selected limit states. 4. conclusions with respect to the situation that current version of eurocode 7 practically does not cover the design of reinforced soil structures, it is a very positive signal to include these structures into the second generation of eurocode 7. this inclusion means that these structures will finally be designed more or less the same way (by the same philosophy) all around europe and even possibly around the world, where the system of eurocodes is being used more and more. the design of reinforced soil structures will be more easily checkable even across borders. for the czech users of eurocode 7 can be my presence in drafting of the new generation a good signal for keeping most design principles of existing design rules in the czech republic (valid since 1997) especially due to the fact that they are used for 20 years without any problems and are fully in accordance with existing eurocode 7 and surely will be also with second generation. acknowledgements here i would like to thank to the czech geotechnical society for the support in my presence on the preparation of second generation of eurocode 7. references [1] čsn en 1997-1:2006. eurocode 7: design of geotechnical structures – part 1: general rules. [2] čsn p env 1997-1:1996. design of geotechnical structures – part 1: general rules. [3] čsn en 1993-5:2008. eurocode 3: design of steel structures – part 5: piling. [4] pd iso/tr 20432:2007. guidelines for the determination of the long-term strength of geosynthetics for soil reinforcement. [5] w. ernst, s. sohn. deutsche gesellschaft für geotechnik e. v., dggt (2011). recommendations for design and analysis of earth structures using geosynthetic reinforcements – ebgeo. 340 p. berlin. [6] čsn en 14490:2010. execution of special geotechnical works – soil nailing. 39 acta polytechnica ctu proceedings 16:37–39, 2018 1 introduction 2 development in between 2010 and 2015 in soil reinforcement area 3 eurocode 7 development since 2015 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0157 acta polytechnica ctu proceedings 13:157–161, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app methodology for measuring the shear stress of plasterboards ondřej zobala, b, ∗, robert járaa, b, martin hataja, b a czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic b czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: ondrej.zobal@fsv.cvut.cz abstract. plasterboards have to meet many requirements for certification. the mechanical properties of itself plasterboard are one of them. the important mechanical property of the plasterboard is the shear strength perpendicular to the plane of the board. the european standard refers to performing the relevant test for the procedure set out in the wood standard en 789. but is this methodology for plasterboard the most suitable? this paper compares this recommended method with the alternative method according to the american standard astm d5379 with indicating the results of a specific measurement. keywords: plasterboard, shear strength perpendicular to the plane of the board, en 789, astm d5379. 1. introduction the plasterboard is pressed gypsum mass between two paper cartons for dry construction. this board can be used for reconstructions and new buildings for the construction of partitions, floors and ceilings. the plasterboard can protect against noise, moisture and fire [1]. marking, definition of properties and test methods for gypsum boards are prescribed standard en 520 + a1 from the 2010 which, moreover, unifies requirements and designations throughout europe. essential requirements include the declared tensile strength in bending and sag, shear strength, compressive strength, hardness and minimum required values of modulus of elasticity for individual stresses [1–5]. all required mechanical properties for certification are given in tab. 1. standard en 520 + a1 defines 8 types of plasterboard: (1.) type a – gypsum plasterboard, (2.) type h – plasterboard with reduced water absorption rate, (3.) type e – gypsum sheathing board, (4.) type f – gypsum plasterboard with improved core adhesion at high temperature, (5.) type p – plasterboard, (6.) type d – gypsum plasterboard with controlled density, (7.) type r – gypsum plasterboard with enhanced strength, (8.) type i – gypsum plasterboard with enhanced surface hardness. this paper addresses the determination of shear strength perpendicular to the plane of the board. european standards din 18180 from 2014 and ead 07001-01-0504 from 2016 refer to the shear strength test perpendicular to the plane of the board to en 789, which is, however, focused to wood-based boards [6– 8]. it has been found that the shear strength test perpendicular to the plane of the board according to en 789 is performed very rare in test institutes in the czech republic (two places) and only one test facility has been discovered on the university campus. for these two reasons, we decided to try a different and more accessible way of testing. the us standard astm d5379 for testing the shear properties of composite materials was finally chosen as an alternative shear strength test perpendicular to the plane of the board [9]. the paper compares results and measurement procedures these two standards. 2. methodology of measurement 2.1. standard en 789 the test specimen had the dimensions of 450×900×thickness of board [mm] and each board was cut one in the longitudinal direction and one in the transverse direction. we needed wooden or steel guide parts for fixing the test specimen. the guide parts had to meet the criteria (minimum dimensions, same humidity as the test specimen, minimum strength and rigidity). the connection of the guide parts and test specimen had to be done with a suitable adhesive. the static scheme is shown in fig. 1 and fig. 2 shows a load arrangement for a shear strength test perpendicular to the plane of the board. 157 http://dx.doi.org/10.14311/app.2017.13.0157 http://ojs.cvut.cz/ojs/index.php/app o. zobal, r. jára, m. hataj acta polytechnica ctu proceedings mechanical properties designa-tion dir. [°] nr. of tests spec. width [mm] spec. length [mm] bending strength perpendicular to the plane of the board – in the longitudinal direction fm,0,k 0 3 400 300 bending strength perpendicular to the plane of the board – in the transverse direction fm,90,k 90 3 400 300 modulus of elasticity in bending perpendicular to the plane of the board – longitudinally em,0,mean 0 modulus of elasticity in bending perpendicular to the plane of the board – transversely em,90,mean 90 shear strength in the plane of the board – longitudinally fv,0,k 0 3 500 600 shear strength in the plane of the board – transversely fv,90,k 90 3 500 600 modulus of elasticity in shear in the plane of the board – longitudinally gv,0,mean 0 modulus of elasticity in shear in the plane of the board – transversely gv,90,mean 90 compressive strength perpendicular to the plane of the board – longitudinally fc,0,k 0 3 400 400 compressive strength perpendicular to the plane of the board – transversely fc,90,k 90 3 400 400 bending strength in the plane of the board – in the longitudinal direction fm,0,k 0 3 400 300 bending strength in the plane of the board – in the transverse direction fm,90,k 90 3 400 300 modulus of elasticity in bending in the plane of the board – longitudinally em,0,mean 0 6 6 modulus of elasticity in bending in the plane of the board – transversely em,90,mean 90 compressive strength in the plane of the board – longitudinally fc,0,k 0 3 400 400 compressive strength in the plane of the board – transversely fc,90,k 90 3 400 400 modulus of elasticity in compressive in the plane of the board – longitudinally ec,0,mean 0 modulus of elasticity in compressive in the plane of the board – transversely ec,90,mean 90 tensile strength in the plane of the board – longitudinally ft,0,k 0 3 50 350 tensile strength in the plane of the board – transversely ft,90,k 90 3 50 350 modulus of elasticity in tensile in the plane of the board – longitudinally et,0,mean 0 modulus of elasticity in tensile in the plane of the board – transversely et,90,mean 90 shear strength core + carton fshear,max 6 600 170 determination of surface hardness of the board 6 300 400 table 1. comparison of test of the shear strength perpendicular to the plane of the board according to the standards en 789 and astm d5379. 158 vol. 13/2017 methodology for measuring the shear stress of plasterboards figure 1. static scheme of the test of shear strength perpendicular to the plane of the board [6]. figure 2. the load arrangement for a shear strength test perpendicular to the plane of the board according standard en 789 (dimensions are in mm) [8]. all test specimens were loaded with incremental strain at a constant rate of 2 mm/min until failure. the shear strength perpendicular to the plane of the board was given by the following equation 1: fv = fmax l.t , (1) where fv is shear strength perpendicular to the plane of the plate [mpa], fmax is the maximum load at failure [n], l is the length of the test specimen measured along the centerline of the shear area [mm], t is the average thickness of the test specimen measured at two points along centerline the shear area [mm]. 2.2. standard astm d5379 the test specimen had the dimensions of 19.1 × 76.2× thickness of board [mm] and it was necessary to cut the specimens in the longitudinal and transverse directions from each board. no guide parts were needed to fasten the test specimen. loading scheme was the same as for the standard en 789 (fig. 1). the loading arrangement with the shape and size of the test specimen for the shear strength test perpendicular to the plane of the board according to astm d5379 is shown in fig. 3. the strength evaluation was performed according to eq. 1. figure 3. the load arrangement for a shear strength test perpendicular to the plane of the board with test specimen according standard astm d5379 (dimensions are in mm) [10]. 3. results of measurement the number of test specimens according en 789 was 6 pieces (3 different boards, 1× longitudinal and 1× transverse direction from each board). the number of test specimens according astm d5379 was 30 pieces (3 different boards, 5× longitudinal and 5× transverse direction from each board). fig. 4 shows a shear strength test perpendicular to the plane of the board according to astm d5379. the characteristic value of shear strength according to en 789 was in the longitudinal and transverse direction same 2.8 mpa. the shear strength according to astm d5379 had characteristic value in the longitudinal direction of 3.1 mpa and in the transverse direction of 2.7 mpa. the differences of values of shear strength with the standard deviation results under 10% were minimal. 159 o. zobal, r. jára, m. hataj acta polytechnica ctu proceedings standard size of the testspecimen [mm] difference of characteristic values [% ] note en 789 600×900×thickness fv,0,k = 10.7 problem with fastening test specimen astm d5379 19.1×76.2×thickness fv,90,k = 3.5 table 2. comparison of test of the shear strength perpendicular to the plane of the board according to the standards en 789 and astm d5379. figure 4. test of shear strength perpendicular to the plate plane according to astm d5379. figure 5. characteristic values of shear strength perpendicular to the plane of the plate according to the standards en 789 and astm d5379. 4. conclusions as is shown in tab. 2, values of shear strength measurements according to en 789 and astm d5379 are very similar. the en 789 standard is determined primarily for wood-based materials. the test specimens according to en 789 are very large and there is a problem with their attachment and ensuring stability during the test. a numerical model will be created to comparing the results according to en 789 and astm d5379 on the basis of the obtained test parameters in the next phase of our work. acknowledgements this work has been supported by the faculty of civil engineering, czech technical university in prague (sgs project no. sgs16/201/ohk1/3t/11) and by the ministry of education, youth and sports within national sustainability programme i, project no. lo1605. references [1] gypsum unique properties in buildings. http: //www.eurogypsum.org/about-gypsum/properties/. [2] plasterboards according to european standards. http://imaterialy.dumabyt.cz/rubriky/ drevene-a-montovane-konstrukce/ sadrokartonove-desky-podle-evropskych-norem_ 101335.html. [3] cni. gypsum plasterboards – definitions, requirements and test methods. first printing. prague, 2010. [4] d. supik. analysis of selected mechanical properties of construction materials for wood building. thesis statement. mendel university in brno, 2009. [5] c. petrone, g. magliulo, g. manfredi. mechanical properties of plasterboards: experimental tests and statistical analysis. journal of materials in civil engineering 28(11):04016129, 2016. [6] din. gypsum plasterboards – definitions, requirements and test methods. first printing. berlin, 2014. [7] ead. gypsum plasterboard for load-bearing applications. first printing. eota, 2016. [8] cni. timber structures – test methods – determination of mechanical properties of wood based panels. first printing. prague, 2005. [9] d. astm. 5379:“standard test method for shear properties of composite materials by the v-notched beam method”. annual book of astm standards 100:235–47, 2012. 160 http://www.eurogypsum.org/about-gypsum/properties/ http://www.eurogypsum.org/about-gypsum/properties/ http://imaterialy.dumabyt.cz/rubriky/drevene-a-montovane-konstrukce/sadrokartonove-desky-podle-evropskych-norem_101335.html http://imaterialy.dumabyt.cz/rubriky/drevene-a-montovane-konstrukce/sadrokartonove-desky-podle-evropskych-norem_101335.html http://imaterialy.dumabyt.cz/rubriky/drevene-a-montovane-konstrukce/sadrokartonove-desky-podle-evropskych-norem_101335.html http://imaterialy.dumabyt.cz/rubriky/drevene-a-montovane-konstrukce/sadrokartonove-desky-podle-evropskych-norem_101335.html vol. 13/2017 methodology for measuring the shear stress of plasterboards [10] shear test fixture astm d5379. http://www.grip.de/pdf/ths600_datasheet.pdf. 161 http://www.grip.de/pdf/ths600_datasheet.pdf acta polytechnica ctu proceedings 13:157–160, 2017 1 introduction 2 methodology of measurement 2.1 standard en 789 2.2 standard astm d5379 3 results of measurement 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0074 acta polytechnica ctu proceedings 11:74–79, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app high speed railway lines – future part of czech railway network? lukáš týfaa, b, ∗, david vodákb a department of logistics and management of transport, ctu in prague, faculty of transportation sciences, prague, czech republic b department of transportation systems, ctu in prague, faculty of transportation sciences, prague, czech republic ∗ corresponding author: tyfa@fd.cvut.cz abstract. the paper first describes high speed rail generally and explains the relationship between high speed and conventional railway networks (according to the vehicle types in operation on the network). the core of the paper is comprised of the methodology for choosing the best route for a railway line and its application to the high speed railway connection praha – brno. the algorithm used assumes the existence of more route proposals, which could be different in terms of the operational conception, line routing or types of vehicles used. the optimal variant is the one with the lowest daily cost, which includes infrastructure and vehicle costs; investment and operational costs. the results from applying this model confirmed the assumption, that a dedicated high speed railway line, only for high speed trains, has the same or lower investment costs than a line for both high speed and conventional trains. furthermore, a dedicated high line also has a lower cost for infrastructure maintenance but a higher cost for buying high speed multiple units. keywords: high speed rail, railway network, investment costs, operational costs. 1. railway network every transportation system has its advantages and disadvantages. this means, that there is a perfect match for every single transport demand (from the sum of all costs – externalities included). in general, railway transportation appears to be the perfect transport system for the backbone of urban and intra-urban transport, regional transport and long-distance continental transport. it also works perfectly in freight transport, as a part of combined transport, or regular transportation between branches. currently railway transportation in the czech republic cannot work optimally because [1]: • conventional mainlines, with standard parameters (from the czech point of view – railway lines with more tracks, maximal speed 120-160 km/h and electric traction), are almost out of capacity and routes in graphical time table cannot be assigned appropriately • some segments have high transport potential due to their route, but their technical and operational parameters (low capacity, low speed, low load class, often speed restrictions) make them unattractive for carriers (transport operators) • on some routes the traffic demand is not satisfied at all or it is satisfied by another mode of transportation, despite the fact, that these routes have perfect conditions for railway transportation • on the other hand, there are some effectively obsolete sections of railway line without any current use and without hope for future transport potential one possible solution, which could improve this situation, is the building of a high speed railway system [1]. the high speed railway system and the conventional railway system are both parts of the general railway system – see directive eu about interoperability of railway system [2]. high speed railway lines include [2]: newly built lines for speeds up to 250 km/h, upgraded conventional lines for speeds of approximately 200 km/h and upgraded conventional lines with special attributes. new high speed railway line is an adhesion railway line (usually track gauge 1,435 mm and electric traction), which has sections with a maximal speed of 250 km/h, and which are long enough so vehicles can reach this speed. combinations of these sections and upgraded sections (usually for the speeds up to 200 km/h) create the high speed railway network. the complexity of railway transportation is mainly caused by strong connectivity between the subsystems, the different life-cycle of the subsystems and the long time period between launch of a project and launch of its regular operation. changes in the operational conception (train traffic schedule) could be made in days and usually lasts one year. it usually takes several years to prepare a railway vehicle for operation and its servisable life is usually between twenty and thirty years. the building of new railway infrastructure usually takes a period of between ten and twenty years (including design and all administrative procedures) and its period of operation could be (with the proper maintenance) a hundred years 74 http://dx.doi.org/10.14311/app.2017.11.0074 http://ojs.cvut.cz/ojs/index.php/app vol. 11/2017 high speed railway lines or so. the situation is similar for high speed railway systems, but due to the higher speed, the connection between subsystems is even stronger. 2. typology of the connection between high speed and conventional network according to the strength of connection between the high speed and conventional railway system, we may establish four basic approaches [1]: (1.) completely segregated railway systems high speed and conventional lines networks are completely separated and independent in this case. conventional trains are using only the conventional lines and high speed trains are using only high speed lines. the advantage of this approach is that we can adjust the parameters of the line to meet the requirements of only high speed trains which usually leads to a smaller radius of curves, higher longitudinal gradients and lower investment costs. the main disadvantage is the impossibility of profiting from connection between the networks and the necessity of a really strong transport flow (passengers and so trains) to make this kind off network economically effective. we can find this kind of network in japan where the main reason for existence of this system is different track gauge on both railway networks. (2.) conventional trains are operating on the high speed railway network in this case, the conventional and high speed networks are connected, but only conventional trains (freight or passenger) are using both networks. conventional trains have to be modified for the operations on the high speed railway lines, especially for the safe passing of high speed trains. the main advantages of this approach are: the possibility of bypassing sections where demand exceeds capacity on the conventional network, the speeding up of conventional trains and/or increasing the usage of high speed railway lines. the disadvantage of this approach is the necessity of adjusting the parameters of the high speed lines to conventional vehicles, which have lower transport performance and lower maximum vehicle speed. in order to adjust the specifications we have to increase the radius of curves and reduce longitudinal gradients which results in higher investment costs. this approach also increases the wear on railway infrastructure (especially rails), because the speeds of the operated trains are different. the higher the difference between speeds is, the more strained the superstructure is. at the same time, the capacity of the high speed line is decreased, because parallel graphical timetables cannot be designed due to the difference in the speeds of trains which results in the need to build passing points, which allow trains with different speeds to pass each other. an example of this solution can be seen in spain, where there are different track gauges on conventional and high speed networks, but where some of the conventional trains are also operating on the high speed lines. these vehicles can change their wheel gauge thanks to the special construction on undercarriage and superstructure. (3.) high speed trains are operating on conventional network in this case the conventional and high speed networks are connected, but only the high speed trains are using both networks; they use the high speed line for the most of train line length. this solution allows high speed trains to get closer to the place of traffic demand without the need for a high speed line. on the other hand, the operation of high speed multiple units on the conventional line is more expensive than operating just conventional trains. the difference in costs is not big, but it still exists – mainly because of higher performance, higher safety and reliability requirements, and better parameters of high speed multiple units (higher mechanical body resistance to air pressure – train passing, running in tunnels). another disadvantage is the higher probability of accidents on the conventional network, which can lead to transition of delay from the conventional network to the high speed network. an example of this solution could be found in france, where high speed trans operate on conventional lines in order to get closer to city centres or to reach destinations, which are not yet connected to high speed network. (4.) high speed and conventional trains are both operating on both networks in this case the conventional and high speed networks are connected and both types of trains are using both networks. this solution uses the maximal potential of connection of both networks. it also brings the possibility of building the high speed railway line in stages, so that non-related segments of the high speed railway line can be built in order to fix the worst issues on the conventional network. the disadvantages are: the higher probability of the accidents on the conventional network, which can lead to a delay in the transition from the conventional network to the high speed network, complicated train traffic schedule designs and stricter the need for higher specifications of the high speed line. this solution also results in the less efficient use of high speed multiple units, because they cannot reach their full potential on the conventional network. an example of this approach can be found on many newly built and upgraded high speed railway lines in germany. the traffic consists not only of the high speed multiple units and regional passenger trains but also of the freight trains (mostly in the night) – see below. 75 lukáš týfa, david vodák acta polytechnica ctu proceedings 3. freight trains on high speed lines there are two basic approaches to freight operation on high speed lines. those approaches are distinguished according to the parameters of freight trains (length, weight, weight on axle, performance). in the first case, the freight train has the same parameters as the passenger train; the difference is only in the transported goods (mail, goods on pallets, special small containers). in the second case the freight train consists of classic freight carriages. in the first case, the freight trains can be treated as passenger trains. in the second case, we have to make some special safety measures to allow for the coexistence of the freight and passenger trains, which usually means, to the determination of different time periods for passenger trains and for freight trains. 4. general benefits of high speed railway network • quantitative – high speed railway lines increase the overall capacity of the railway network but it is important connecting new lines to the nodes must not cause traffic restrictions or instability on the existing conventional network • qualitative – better transportation by cutting the travel times between main cities • economical for railway carriers (transportation operators) – high speed multiple units are going to run more distance per day and more distance on high speed lines • decreasing intensity of road transportation by moving the transport flows to the railway with significant benefits in terms of safety and the environment • establishing connections between trains in periodical timetable • secondary effects – increasing share of electrical traction (and so production decreasing of carbon dioxide and harmful exhalation, possibility of independence from hydrocarbon fuels) 5. methodology of project comparison and its application the main goal of this paper was to introduce an algorithm for choosing the best route for a railway line and its application to the high speed railway connection praha – brno. the algorithm assumes the existence of more route proposals, which could be different in operational conception, line routing or types of vehicles. the optimal variant (according to the algorithm) is the one with the lowest daily cost, which includes infrastructure and vehicle costs and investment and operational costs. the algorithm does not take into account the dividing of the railway transportation among subjects – railway transportation is assumed as one unit. this unit consists of owning the railway, operating the railway and operating the railway transportation. the methodology is showed in the six connected tables. the first table, "operation", shows operated lines on the line; then number of trainsets (multiple units) and the daily time costs of trains in operation are calculated. the second table, "trainsets", shows the trainsets (multiple units) operated on the line and its results are the daily costs for buying trainsets. table iii. "unitary infrastructure investment costs", shows aggregated investment costs for the infrastructure – it is used for drawing up a budget. table iv. "infrastructure and operation on it" is designed to calculate the infrastructure investment and train operational costs according to energy consumption. table v. "track sections with constant operation" calculates the infrastructure and trains operating costs and the infrastructure investment costs. table vi. "recapitulation" shows the unitary daily costs from the previous tables and also divides daily the cost between the subjects on the railway (infrastructure manager, carriers – railway operators, infrastructure owner). to demonstrate the algorithm, two variants of high speed railway line praha – brno were evaluated. the first variant is from the study designed by sudop praha ag in 2010 [3], which is called study 2010. for the purposes of demonstration, variant h4 from the study 2010 was chosen. the second variant is from the study lt 2012, which is study 2010 with alterations from lukáš týfa, one of the authors of this paper [1]. study 2010 assumes following operational conception – the high speed railway network will be used both by high speed and conventional trains, high speed trains will be mainly operated on the high speed railway lines and only temporarily on conventional lines, conventional trains will use connections between both networks to enter the high speed lines and use them for the part of their journey. variant lt 2012 assumes the same horizontal alignment as study 2010, but makes changes in the vertical alignment – maximal proposed longitudinal gradient value is 35‰. this is a consequence of the change in the operational conception – line segment benešov – brno will be used only by high speed multiple units, those which can be easily operated in such a longitudinal gradient. a change in the operational conception could also cause a lowering of the minimal curve radius, but the author of variant lt 2012 thinks, that this value was defined too benevolent in study 2010, so no changes were made in this value [4, 5]. both variants have same maximal proposed speed – 350 km/h, although regular operations with a speed 350 km/h do not seem very effective (only a minor cutting in travel times , but a major increase of the operational costs). the main design parameters (track geometry characteristic) of study 2010 and lt 2012 are displayed in table 1. 76 vol. 11/2017 high speed railway lines study/variant study 2010 lt 2012 parameter var. v7 var. h4 line speed limit (max. train speed) v [km/h] 300 330 350 350 350 350 300 350 standard speed of the slowest trains vmin [km/h] 140 160 160 200 200 200 300 350 max. cant dmax [mm] 180 173 163 157 65 150 180 170 max. cant deficiency imax [mm] 150 100 100 80 80 130 150 80 max. cant excess emax [mm] 109 109 109 80 18 80 -* -* min. radius of curve rmin [m] 3 220 4 720 5 500 6 100 10 000 5170 3 220 5 790 type of transition curve and cant transition nonlinear nonlinear nonlinear non-linear non-linear nonlinear nonlinear nonlinear max. longitudinal gradient smax [%] 20 20 20 20 20 20 35 35 track construction slab slab slab ballasted ballasted slab slab slab max. length of tunnel by tsi lt ,max [km] 5 5 5 5 5 5 20 20 * cant excess evaluation is considered only in large longitudinal gradients when high speed train cannot reach line speed table 1. main design parameters of study 2010 and lt 2012. evaluation of variants study 2010 and lt 2012 shows that the daily costs of study 2010 are about a million czk higher (see table 2 and figure 1). the daily costs for buying trainsets are 3% higher for variant lt 2012 – it considers only the high speed multiple units operation with the speed 300 km/h or 350 km/h. both variants have the same investment infrastructure costs. variant lt 2012 assumes slab track, which is more expensive than ballasted track, but this variant saves money on building tunnels and bridges. thanks to the maximal longitudinal gradient 35‰ (in comparison with 20‰ in study 2010) these structures are less frequent in this variant lt 2012. slab track also resulted in a 7% cost decrease, because maintenance costs are reduced. a negligible part of the costs includes: trainsets cleaning, overheads and investment to estates (estates costs are divided into long time). the different composition of costs for both of the variants causes different cost ratios between the infrastructure manager and railway transportation operator. in the study 2010, the railway transportation operator spends 60% of overall costs. in the variant lt 2012, the railway transportation operator spends 67% of overall costs. this distribution also has an impact on infrastructure manager’s daily costs for the unit of line length and the railway transportation operator daily costs, because the costs are 22% lower for variant lt 2012 and, because the costs are 14% higher for variant lt 2012. the difference in infrastructure investment costs are approximately three billions czk. 6. conclusion there is no doubt that the building and operation of a high speed railway system is one of the key steps to improve railway transportation functionality [6]. we can testify this not only in the states of western europe or japan, but also in morocco and saudi arabia. there are many different approaches to the relationship between current conventional networks and newly built high speed networks. these approaches are based on demographical structure, terrain configuration, conventional network parameters and the overall goals for railway transportation [7]. the choice of approach affects the relationship between the infrastructure and the train parameters. the different approaches to conventional/high speed train integration were described in the first part of this paper and applied on the planned high speed railway line praha – brno, which should create the best transport route in the czech republic. the best variant was chosen according to the total intern costs – investment and operational. for the purpose of calculation unit costs were used t. the values of these costs are different, according to what sources and assumptions are made and are subject to debate [8–10], but every cost which was used, is considered as representative and approximately the correct value. not only was a theoretical model made, but a fully functional and editable model was created for the final calculation of costs. the application of the model has confirmed the assumption that a high speed railway line, for use exclusively for high speed trains has the same or lower 77 lukáš týfa, david vodák acta polytechnica ctu proceedings study – variant study 2010 – var.h4 lt 2012 item costs per day costs per day costs per day [czk/day] [%] [czk/day] [%] hs train operation (time costs) 263 571 2 258 844 2 train-sets acquisition 1 365 766 10 1 713 718 13 investment to line-section infrastructure 3 086 366 22 2 933 059 22 investment to estates 11 707 0 11 798 0 electric power for trains 3 238 855 23 3 176 694 24 hs line infrastructure maintenance 2 521 874 18 1 464 062 11 hs train-sets maintenance 3 542 821 25 3 473 389 26 hs train-sets cleaning and overhead costs 162 182 1 145 023 1 sum 14 193 142 13 176 587 costs per day of railway transportation subjects: infrastructure manager 40% 33% carriers 60% 67% infrastructure owner 0% 0% infrastructure investment costs: total investment costs of infrastructure construction [mil. kč] 96 348 93 162 total costs of purchase estates [mil. kč] 427 431 table 2. costs evaluation of variants. figure 1. cost evaluation of variants. 78 vol. 11/2017 high speed railway lines investment costs than line for both high speed and conventional trains (the higher cost for slab track are compensated for by reduced number/length of the tunnels and bridges). a high speed line also has lower cost for infrastructure maintenance but a higher cost associated with buying only high speed multiple units. the costs for a line exclusively for high speed trains (high speed multiple units) is lower in the absolute terms in comparison to line for both high speed and conventional trains. the selection of the optimal variant seems to be clear, but minimization of total internal costs does not have to be the only important parameter. for example, we can decide only according to the total infrastructure manager investment costs. this shows strong connectivity between the railway transportation subsystems and the importance of continuous coordinated planning and design of the whole railway system. after we overcome all of the obstacles, society, as a whole, will be rewarded with a safe, environmentally friendly and highly effective railway system. this is why it is necessary to examine possible parameters (operational conception, line parameters, safety devices, train sets) of the new lines, before we even assign routing studies to the designers. furthermore, we cannot omit an evaluation of the time period between launching the project and launching its regular operation for all subsystems and their life cycles. it is possible to conclude that a high speed railway system in the czech republic should consist of several line types (or line segments). the parameters of these types should be based on the anticipated types of trainsets (multiple units), which are going to use the system. some segments should have parameters that fit both conventional trains and high speed trains, while others should be designed only for high speed trains. the individual types of lines should therefore be designed with different parameters. decisions on the types of high speed lines segments will significantly affect current and future conditions of the conventional railway network, which, once again, shows us strong connectivity between different parts of railway system. references [1] l. týfa. service of high speed railway systems in region [associate professor lecture]. praha: czech technical university in prague, 2013. isbn 978-80-01-05199-3. [2] l. týfa. nejnovější trendy v oblasti infrastruktury vysokorychlostních tratí. in odborná konference vysokorychlostní železniční doprava ve světě a v české republice [cd-rom]. praha: sudop praha a.s., 2007. [3] directive eu no. 2008/57/ec on the interoperability of the rail system within the community. [4] čsn en 13803-1:2010. railway applications – track – track alignment design parameters – track gauges 1435 mm and wider – part 1: plain line, 2010. [5] j. tvrdík. podmienky pre výstavbu vrt praha – brno. in 7. fórum koľajovej dopravy. bratislava: spoločnosť pskd – prevádzka a stavby koľajovej dopravy, 2011. isbn 978-80-88973-68-3. [6] j. ilík, j. kušnír, r. čech. budoucnost české železnice v dokumentech evropské dopravní politiky. železniční magazín. praha: m-presse plus, 2012, roč 18, č 1. issn 1212-1851. [7] uic. high speed rail: fast track to sustainable mobility. paris: uic, high speed department. 2nd ed. updated. isbn 2-7461-1452-6. [8] j. pohl. rychlá železniční osobní doprava: díl dvacátý: náklady a výnosy poprvé. železniční magazín. praha: m-presse plus, 2010, roč 16, č 1. issn 1212-1851. [9] j. pohl. rychlá železniční osobní doprava: díl dvacátý první: náklady a výnosy podruhé. železniční magazín. praha: m-presse plus, 2010, roč 16, č 2. issn 1212-1851. [10] i. c. gmbh, inrets. traffic and profitability for a western european high speed train network, 1994. 79 acta polytechnica ctu proceedings 11:74–79, 2017 1 railway network 2 typology of the connection between high speed and conventional network 3 freight trains on high speed lines 4 general benefits of high speed railway network 5 methodology of project comparison and its application 6 conclusion references 21 acta polytechnica ctu proceedings 2(1): 21–25, 2015 21 doi: 10.14311/app.2015.02.0021 the ritter-kolb catalogue and its impact on research into cvs, lmxbs and related objects h. ritter1, u. kolb2 1max–planck–institut für astrophysik, d–85741 garching, germany 2department of physical sciences, the open university, milton keynes mk7 6aa, united kingdom corresponding author: hsr@mpa-garching.mpg.de abstract in the first part of this paper, the ritter-kolb catalogue (rkcat for short), its history and a few examples of its application to research in the field of cataclysmic variables are briefly described. in a second part we look forward to possible future applications of rkcat for the study of cataclysmic variables, low-mass x-ray binaries and related objects. last but not least we also briefly comment on the future of the rkcat service itself. keywords: cataclysmic variables cataclysmic binaries low-mass x-ray binaries post-common envelope binaries catalogues. 1 introduction thirty years after the first public circulation of the catalogue of cataclysmic binaries, low-mass x-ray binaries and related objects, the so-called ritter catalogue or rcat for short (ritter 1982) it is perhaps a good time to look back at its history and some of its applications to research in the field of cataclysmic variables (cvs), lowmass x-ray binaries (lmxbs) and related objects, and at the same time to also look ahead to possible future applications of what later became to be known as the ritter-kolb catalogue (ritter & kolb 2003), hereafter rkcat for short. but before discussing the history and the future of rkcat let us briefly summarize for those who are not familiar with rkcat what rkcat actually is, i.e. its purpose, which information it does provide and how it can be accessed. 1.1 rkcat: what is it? ever since it was conceived rcat (up to 1998) and later rkcat (from 1998 onwards) has provided a comprehensive compilation of detached and semi-detached postcommon envelope binaries with known (or suspected) orbital periods. the restriction to systems with “known orbital period” is important for two reasons: first, knowing the orbital period of a binary not only provides important information about the object in question, but it is usually also a sure sign for the object having undergone serious examination to some degree, and that more is known about it than just its position on the sky, its apparent magnitude, and that it may be variable. second, restriction to systems with known orbital period also considerably limits the number of objects to be dealt with which is anyway growing at an alarming rate (see below). among the semi-detached systems we then have cvs (with at least one white dwarf (wd) component) and lmxbs (with either a neutron star or a black hole primary and usually a non-degenerate companion, but low-mass wds do also occur). among the detached systems, in rcat/rkcat referred to as “related objects”, we have “white dwarf”-red dwarf binaries and “white dwarf”-”white dwarf” binaries (double degenerates), where a “white dwarf” is either a genuine wd or an immediate precursor of a wd such as a sdbor sdostar. not included in rcat/rkcat among the detached systems are binary radio pulsars although some of them do in fact match the criteria for inclusion among the “related objects”. the reason for excluding binary radio pulsars is because they are well documented elsewhere (manchester et al. 2005).1 1.2 rkcat: what does it contain? rkcat provides information about each object in tabular form, a list of references to the most relevant recent literature, a list of references to published finding charts and a comprehensive cross reference list of alias object designations. in the tables (separate for cvs, lmxbs and related objects) we provide object designations, the position on the sky (α,δ), information about the object 1the data are available electronically at http://www.atnf.csiro.au/research/pulsar/psrcat. 21 http://dx.doi.org/10.14311/app.2015.02.0021 h. ritter, u. kolb type/subtype, apparent magnitudes, periods (above all the orbital period, but also the superhump period and the spin period of the wd (in cvs), or of the neutron star (in lmxbs)), the spectral types of the components as well as other orbital and stellar parameters (orbital inclination, mass ratio, masses, radii) to the extent that these quantities are known. the purpose of the cross reference list of alias object designations (appropriately called “whoswho”) is to help avoid confusion about object idendity. many objects have a long history during which they have been given many different names by different authors. thus confusion about an object’s idendity can simply arise because earlier authors can of course not refer to later object designations. but also going backwards in the literature does not always reveal even the more common alias names. the “whoswho” list is an attempt to remedy that situation and to nail down an object’s idendity in addition to its identification on a finding chart. the most recent release of rkcat at the time of writing (ritter & kolb 2003, update rkcat7.20 as of 30 june 2013) contains information about 1093 cvs, 104 lmxbs and 483 related objects, i.e. of 1680 objects in total. 1.3 rkcat: how can it be accessed? whereas earlier editions of rcat/rkcat appeared in print (ritter, 1982, 1983, 1984, 1987, 1990; ritter & kolb 1995, 1998) the current (7th) edition (ritter & kolb 2003) is available in electronic form only from websites at the open university (ou) at http://physics.open.ac.uk/rkcat and at the max-planck-institut für astrophysik (mpa) at http://www.mpa-garching.mpg.de/rkcat, and in slightly different form but with the same content from the strasbourg astronomical data center (cds) at http://cdsarc.u-strasbg.fr/viz-bin/cat?b/cb. the main difference between the versions provided by the ou/mpa websites and the cds website is that the former provide the tables in a two lines per object format with delimeters which allows for easy visual inspection of the tables on a computer screen, whereas the cds tables come in a one line per object format which (even without the delimiters) is much too wide for that purpose. rather they are designed to be processed electronically. many more copies of rkcat are available from various other websites. but because the contents of these copies is beyond our control we do not take responsibility for the content of any other than the abovementioned sources. 2 a brief history of rcat/rkcat compiling data on cvs for what was later to become rcat actually began already in 1973, i.e. 40 years ago, as part of the first authors’s (hereafter hr) phd project (ritter 1977). although at that time there was no intention whatsoever to produce a catalogue, compiling data on cvs, lmxbs, and related objects continued also beyond the end of the phd project, and it was around 1978/79 that rumors about the existence of this compilation began to spread among colleagues, some of whom did not hesitate to ask for copies. dutifully, hand-written copies of the compilation were produced and circulated among a few priviliged individuals. in that context, one must realize that at that time neither the internet nor email did exist, and electronic storage of data on computers was unavailable the storage media of choice were either hand-written files or punched cards and consequently, there was no easy way of editing and updating such a data base. bogdan paczynski was among the first to whom a copy was sent upon his request. and based on the distribution of the orbital periods of the 43 cvs known at that time he must have realized the sharp drop of the number of cvs shortward of an orbital period of ∼ 80min, and not long after that went on explaining the physical significance of the minimum orbital period of cvs in the context of cv evolution (paczynski 1981; paczynski & sienkiewicz 1981, 1983). this was probably the first direct application of the immediate precursor of rcat to cv research. around 1981 personal storage accounts and screen editors on computers became finally available. therefore, the data of the cv compilation were migrated to permanent electronic storage and editing. from that time on it not only became much easier to update the compilation but also to produce printed copies of it. as a consequence, so to speak as a test, soon after having gone electronic, the first and second edition of rcat were produced and circulated in preprint form among a relatively large number (∼ 200) of colleagues and libraries (ritter 1982, 1983). the positive response to the catalogue then lead to the publication of the 3rd edition printed in a refereed journal (ritter 1984). with data on cvs thus becoming widely available it took not long before another of the remarkable properties of the cv period distribution, i.e. the period gap, was addressed and an explanation for it in terms of cv evolution was given (spruit & ritter 1983; rappaport, vebunt & joss 1983). around 1987, when the 4th edition of rcat (ritter 1987) was published it was also realized, not least thanks to x-ray observations with exosat, that the distribution of the orbital periods of the synchronized and non-synchronized magnetic cvs, i.e. of the am her stars and the intermediate polars, are significantly dif22 the ritter-kolb catalogue and its impact on research into cvs, lmxbs and related objects ferent and that this could be accounted for in terms of standard cv evolution by the fact that on average the mass transfer rate in cvs decreases with orbital period (hameury et al. 1987). with the publication of the 5th edition of rcat (ritter 1990) another interesting property of the period distribution of cvs became apparent, namely the peculiar distribution of the orbital periods of am her stars below the period gap with 6 am her stars piling up within a very narrow period interval of only 2 min, the so-called period spike (see fig. 1 in ritter & kolb 1992). in that paper we had also given a detailed analysis of the period spike and explained its position and width as a natural consequence of cv evolution in the context of the disrupted magnetic braking hypothesis. although at that time the am her period spike was considered to be statistically highly significant and thus worth studying, now, more that 20 years later and with many more systems known, we must admit that there is no such spike and that we were just fooled by a statistical fluke. with the 6th edition of the catalogue (ritter & kolb 1998) two major changes occurred. first, because the amount of work connected with updating the catalogue had increased so much that additional support, in particular for servicing the newly established catalogue websites, was needed, ulrich kolb (uk) stepped in, and so rcat became rkcat. since then uk has been mainly responsible for the maintenance of our websites. second, beginning with the 6th edition the data became available in electronic form only, and no more than a brief introduction to the catalogue has appeared in print. finally since 2003, the 7th edition of rkcat (ritter & kolb 2003) has become available, the data again in electronic form only. the catalogue websites of the 7th edition are now updated regularly, currently every 6 months, and at the time of writing the 20th update (release 7.20) is available. the development of rcat/rkcat over the past 35 years is perhaps best illustrated by plotting the number of objects contained in the catalogue as a function of time, as is shown in fig. 1. as is readily seen, the number of objects has been steadily and exponentially increasing with a growth rate of ∼ 9%/year over the past 30 years! 3 some examples of possible future applications of rkcat before going to discuss a few specific possibilities for future applications of rkcat, let us briefly make two rather general comments concerning rkcat and its future use. • the first comment is a warning, namely that based on the criteria according to which objects enter the catalogue, no subset of the objects listed in rkcat represents a statistically well-defined sample. one has to keep this in mind when e.g. comparing results of population synthesis calculations with distribution functions extracted from rkcat. • the second general comment is that by now the number of objects contained in rkcat is larger by a factor 10 to 20 than it was 20 30 years ago, i.e. when the first applications of rcat to problems of cv evolution were made. what this means is that now the contents of rkcat allows for statistically meaningful studies of the properties of various subsets of cvs, lmxbs or related objects, which because of too small numbers of objects had not been possible a decade or two ago. figure 1: number of cvs (open squares), lmxbs (open triangles), and related objects (filled diamonds), as well as the total number of objects (filled circles) contained in rcat/rkcat and its precursors as a function of time. let us now turn to just a few specific possible future applications of rkcat. 1. it has long been known that the average orbital period of age zero post common envelope wd-rd binaries, i.e. of the binary central stars of planetary nebulae and systems where the wd is still so hot that it gives rise to a noticeable reflection effect on the companion, is systematically longer (by about a factor of 2) than the average orbital period of the remaining detached post common envelope wd-rd binaries. the qualitative explanation for this is that post common envelope binaries undergo orbital evolution as a consequence of 23 h. ritter, u. kolb loss of orbital angular momentum. in that context it is important that now rkcat contains a relatively large number of detached wd-rd post common envelope binaries with reasonably accurately known binary parameters. based on these data it should be possible to investigate the angular momentum loss of post common envelope wd-rd binaries in more detail, e.g. its dependence on properties of the rd, i.e. whether it is evolved or fully convective or not, and maybe even quantify it. 2. the period distribution of dwarf novae shows an undisputable dearth of systems in the period range 3h . porb . 4 h which is not predicted by the standard cv evolutionary paradigm, i.e. the disrupted magnetic braking hypothesis, in combination with the disc instability model for dwarf nova outbursts. a satisfactory explanation for this phenomenon is so far lacking. in the framework of the disc instability model this could be accounted for only if the mass transfer rate is decreasing less strongly with orbital period than the critical mass transfer rate for dwarf nova outbursts as a cv evolves from longer to shorter orbital periods. this, in turn, implies that the angular momentum loss time scale must decrease with the orbital period, contrary to what standard prescriptions of magnetic braking, e.g. verbunt & zwaan (1981) or mestel & spruit (1983) predict. the ratio of the number of dwarf novae to novalike systems as a function of orbital period, on the other hand, should allow for an empirical calibration of the angular momentum loss rate as a function of orbital period. 3. as the absence of a clear period gap in the period distribution of am her systems suggests, the strong magnetic field of the wd interferes with the loss of mass and angular momentum via a stellar wind from the companion star with the result that magnetic braking in magnetic cvs is significantly weaker than in non-magnetic cvs. therefore, if below the period gap angular momentum loss via magnetic braking is comparable to that caused by gravitational radiation, as has been argued e.g. by knigge, patterson & baraffe (2011), and if magnetic braking is also weaker in am her stars below the gap, one would expect that also the minimum period of am her stars is systematically shorter than that of non-magnetic systems. therefore, by a careful examination of the period distributions of magnetic and non-magnetic cvs below the gap it should be possible to get more insight into the evolution of cvs below the gap. clearly, the above list of possible future applications of rkcat is not exhaustive. what we hope for is that our colleagues working with rkcat will come up with many more and even more ingenious applications of rkcat for the research in the field of compact binaries than we have just given. 4 the future of rkcat apart from discussing future appliactions of rkcat we think that writing this paper provides us also with an exellent opportunity to briefly address the possible future of rkcat itself. in that context it is important to be aware of the following facts: • the number of objects in rkcat has been growing exponentially at a steady rate of ∼ 9%/yr over the past 30 years (see fig. 1), and so has the amount of work that has been going into rkcat. and there is no reason to expect that this will be much different in the near future. • already now, servicing the catalogue data base is practically a full-time job for hr. • hr is already several years beyond formal retirement age. taken these three points together it is obvious that hr will definitely not be able to provide this service for many more years. if rkact or something akin to it should be continued beyond hr’s definitive retirement from that job then it is clear that in view of the first two points above some changes in how this service is and in what should be provided are unavoidable. the all decisive question to the cv community then is: who is prepared to continue with rkcat and in which form? anybody out there? acknowledgement hr would like to thank franco giovanelli for inviting him to attend this workshop and thus for providing him an opportunity to talk about the past and future of rkcat. references [1] hameury, j.m., king, a.r., lasota, j.p., ritter, h.: 1987, mnras 231, 535. [2] knigge, c., baraffe, i., patterson, j.: 2011, apjs 194, 28. 24 the ritter-kolb catalogue and its impact on research into cvs, lmxbs and related objects [3] manchester, r.n., hobbs, g.b., teoh a., hobbs, m.: 2005, aj 129, 1993. doi:10.1111/j.1745-3933.2006.00242.x [4] mestel, l., spruit, h.c.: 1987, mnras 226, 57. [5] paczynski, b.: 1981, aca 31, 1. doi:10.1017/cbo9781139164245 [6] paczynski, b., sienkiewicz, r.: 1981, apj 248, l27. doi:10.1086/421295 [7] paczynski, b., sienkiewicz, r.: 1983, apj 268, 825. [8] rappaport, s., verbunt, f., joss, p.c.: 1983, apj 275, 713. doi:10.1086/382481 [9] ritter, h.: 1977, phd thesis, univ. hamburg. doi:10.1086/590491 [10] ritter, h.: 1982, catalogue of cataclysmic binaries, low-mass x-ray binaries and related objects, first edition, preprint mpa 22. doi:10.1046/j.1365-8711.1999.02926.x [11] ritter, h.: 1983, catalogue of cataclysmic binaries, low-mass x-ray binaries and related objects, second edition, preprint mpa 51. [12] ritter, h.: 1984, a&as 57, 385. [13] ritter, h.: 1987, a&as 70, 335. [14] ritter, h.: 1990, a&as 85, 1179. doi:10.1046/j.1365-8711.2000.03780.x [15] ritter, h., kolb, u.: 1992, a&a 259, 159. [16] ritter, h., kolb, u.: 1995, in x-ray binaries, w.h.g. lewin, j. van paradijs, e.p.j. van den heuvel (eds.), cambridge astrophysics series 26, cambridge university press, p. 578. [17] ritter, h., kolb, u.: 1998, a&as 129, 83. doi:10.1086/341696 [18] ritter, h., kolb, u.: 2003, a&a 404, 301. [19] spruit, h.c., ritter, h.: 1983, a&a 124, 267. doi:10.1088/0004-637x/721/2/1356 [20] verbunt, f., zwaan, c.: 1981, a&a 100, l7. discussion ashley pagnotta: regarding the future of rkcat, have you considered working with the aavso? their vsx is not nicely curated like rkcat but is a possible future home. hans ritter: up to now i had not approached anybody or any organisation regarding the future of rkcat. this is actually the first time that i have spoken in public about that problem. christian knigge: just a comment on the topics you suggested for further study, which i think are really interesting. regarding pmin for magnetic cvs, boris gaensicke, retha pretorius and i have all looked at this a bit and have not seen the expected difference to non-magnetic cvs. [note added: i think this is mostly unpublished – the only relevant publication i am aware of is pretorius, knigge & schwope 2013, mnras, 432, 570, where fig. 1 shows that the period minimum for magnetic cvs is consistent with that for all cvs, which are dominated by non-magnetic systems. but we don’t really discuss that in the text.] also, in fig. 18 of knigge, baraffe & patterson (2011) we plot the dwarf nova fraction versus porb and find that not only is it zero just above the gap, but going from shorter to longer orbital periods it declines monotonically through the gap, which is very surprising. — 25 http://dx.doi.org/10.1111/j.1745-3933.2006.00242.x http://dx.doi.org/10.1017/cbo9781139164245 http://dx.doi.org/10.1086/421295 http://dx.doi.org/10.1086/382481 http://dx.doi.org/10.1086/590491 http://dx.doi.org/10.1046/j.1365-8711.1999.02926.x http://dx.doi.org/10.1046/j.1365-8711.2000.03780.x http://dx.doi.org/10.1086/341696 http://dx.doi.org/10.1088/0004-637x/721/2/1356 introduction rkcat: what is it? rkcat: what does it contain? rkcat: how can it be accessed? a brief history of rcat/rkcat some examples of possible future applications of rkcat the future of rkcat acta polytechnica ctu proceedings doi:10.14311/app.2018.18.0024 acta polytechnica ctu proceedings 18:24–27, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app influence of printing and loading direction on mechanical response in 3d printed models of human trabecular bone tomáš doktora, ∗, ivana kumpováa, sebastian wrońskib, maciej śniechowskib, jacek tarasiukb, giancarlo fortec, daniel kytýřa a intitute of theoretical and applied mechanics of the cas, prosecká 76, 19000 prague 9, czech republic b agh university of science and technology, aleja adama mickiewicza 30, 30-059 kraków, poland c international clinical research center of st. anne’s university hospital brno, pekařská 53, 656 91 brno, czech republic ∗ corresponding author: doktor@itam.cas.cz abstract. the paper deals with investigation on directional variations of mechanical response in 3d printed models of human trabecular bone. sample of trabecular bone tissue was resected from human donor and 3d model was obtained by x-ray computed tomography. then a series of cubical samples was prepared by additive manufacturing technique and tested by uniaxial compression loading mode. mechanical response was compared in nine different combinations of direction of 3d printing and loading direction. the results show neglectible influence on the deformation response in elastic region (stiffness) and significant changes of the behaviour in plastic region (stress and strain at yield point, strain at full collapse). keywords: additive manufacturing, tomography, deformation response. 1. introduction in the regenerative medicine bone scaffolds serve to support the bone healing process [1, 2]. additive manufacturing provides possibilities for manufacturing of the scaffold structure with tailored properties. moreover, mechanical response and topological properties of additively manufactured stuctures may be investigated and optimised using experimental modeling [3, 4]. however, direction of the 3d printing related to the loading direction may influence the effective mechanical properties and therefore affect the reliability of such analyses. in order to investigate the influence of the printing direction a series of uniaxial loading tests was performed on additively manufactured models of the human trabecular bone. 2. materials and methods 2.1. sample preparation trabecular bone tissue was extracted from a cadaver human donor (male, 52 years, non pathological conditions). the extraction was performed using hollow drill using a low-speed drilling machine in direction coaxial with column femoris. diameter and length of the resected sample were 18 mm and 40 mm, respectively. the sample is depicted in figure 1 the resected tissue was then cleaned in ultrasonic cleaner using detergent solution in order to remove bone marrow. figure 1. trabecular bone sample resected from human caput femoris. scale bar 20 mm. 2.2. tomographic scanning and reconstruction the extracted sample was then tomographically scanned using toratom (unique patented tomographic facility placed in centre of excelence in telc, itam, czech republic) [5]. for the irradiation x-ray tube xwt 160 tchr (x-ray worx, gmbh, germany) was used with tube voltage 90 kv and target current 76 µa. for the acquisition flat panel detector xrd 1622 ap 14 (perkin elmer, usa) was used. resolution of the detector was 2048 × 2048 px. the tomographic setup is depicted in figure 2. to achieve sufficient magnification, the following geometry of the tomographic setup was used: focus-detector distance was 1200 mm and focus-sample distance was 60 mm, which resulted in magnification 20 ×. duration time of the acquisition was 2000 ms per projection and resulting pixel size was 10 µm. 24 http://dx.doi.org/10.14311/app.2018.18.0024 http://ojs.cvut.cz/ojs/index.php/app vol. 18/2018 influence of printing and loading direction in 3d printed models figure 2. toratom scanning setup. acquired tomographic projections were then recontructed using vg studio software tool (volume graphics, gmbh, germany). from the reconstructed matrix a cubical region of interest (roi) with size 1000 × 1000 × 1000 px3 was selected for 3d print. visualisation of the roi is depicted in figure 3. figure 3. selected region of interest in the reconstructed 3d matrix. scale bar 10 mm. 2.3. additive manufacturing subsequently, nine samples were manufactured using 3d printer, three in each group printed along different axes. the samples were prepared at additive manufacturing facility placed at agh university krakow. custom modified device b9creator (b9creations, llc, usa) was used together with base material vorex resin (madesolid, usa) with adition of photoabsorber sudan i (sigma aldrich, usa). cubic samples with edge length 10 mm were prepared (depicted in figure 4). resolution in the printed plane corresponded to pixel size 30×30 µm and the layer thickness was 30 µm as well. as typical thickness of human trabecula in the selected location varies in range 120 to 160 µm [6] at least four voxels across the trabeculae thickness were present in the printed model. hence the 3d printer resolution was sufficient to preserve all features of the trabecular structure without scaling of the model. figure 4. additively manufactured trabecular bone structure. scale bar 10 mm. 2.4. uniaxial compression tests the additively manufactured samples were then subjected to uniaxial compression using a custom loading frame [7]. loading capacity od the loading frame was 2 kn. loading was provided by a stepper motor attached to a harmonic drive and controlled by an inhouse control software based on linuxcnc interface [8]. the setup was instrumented by force transducer (type u9b, hbm, germany) with loading capacity 1 kn. the loading setup is depicted in figure 5. displacement controlled compression tests with loading rate 30 µm · s−1 were performed. final displacement corresponding to overall strain 0.55 was prescribed (entire available travel range of the loading frame was used to ensure description of both elastic region and plastic collapse of the cellular structure). however in several tests the loading was stopped at lower displacement due to the force drop which indicated complete failure of the structure. 2.5. strain measurement loading scene was captured by monochromatic ccd camera avt manta g504b (alied vision gmbh, germany) attached to a bi-telecentric lens tzcr072 (optoengineering, italy) at resolution 2452 × 2056 px at 2 fps. for camera read out in-house acquisition software [9] based on opencv library [10] was used. 25 t. doktor, i. kumpová, s. wroński et al. acta polytechnica ctu proceedings figure 5. custom loading setup for uniaxial compression tests. displacement of the loading supports was tracked by custom dic software tool [11] based on lucas-kanade trancking algorithm [12]. the overall strain was measured optically based on projections of the loading scene using digital image correlation algorithm (dic). due to the nearly transparent nature of base material of the samples the displacement tracking was not successfull at the specimens’ surface and dic was used for tracking of the loading supports only. 3. results from the calculated displacements and force records stress-strain diagrams were constructed and used for comparison among the tested groups. the obtained loading curves are depicted in figures 6, 7 and 8, where the samples were groupped with respect to the loading direction in order to provide comparison between different printing directions. 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0 20 40 60 80 100 0 1 2 3 4 5 e n g . s tr e s s [ m p a ] l o a d in g f o rc e [ n ] eng. strain [-] displacement [mm] printed along x axis printed along y axis printed along z axis figure 6. loading curves of samples loaded along x direction. 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0 20 40 60 80 100 0 1 2 3 4 5 e n g . s tr e s s [ m p a ] l o a d in g f o rc e [ n ] eng. strain [-] displacement [mm] printed along x axis printed along y axis printed along z axis figure 7. loading curves of samples loaded along y direction. 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0 20 40 60 80 100 0 1 2 3 4 5 e n g . s tr e s s [ m p a ] l o a d in g f o rc e [ n ] eng. strain [-] displacement [mm] printed along x axis printed along y axis printed along z axis figure 8. loading curves of samples loaded along z direction. 26 vol. 18/2018 influence of printing and loading direction in 3d printed models 4. conclusions based on the uniaxial loading tests of different sets of 3d printed specimen of human trabecular bone influence of printing and loading direction on deformation behaviour was studied. compared to previous studies on deformation response of trabecular tissue (human femur [13] or rat vertebra [14]) the additively manufactured models exhibit similar behaviour up to the first plastic collapse in terms of continuous linear elastic region with smooth yield region. the mechanical response was compared in terms of effective young’s modulus, yield stress and strain and peak force at failure. stress-strain diagrams of the 3d printed models of trabecular bone were consistent in the groups with each loading direction in terms of stiffness (slope of the linear response). highest peak force was observed in samples loaded along the direction of 3d printing in all three groups. in several samples full collapse of the cellular structure occurred when the loading axis was perpendicular to the printing direction. acknowledgements the research has been supported by the european regional development fund in frame of the projects kompetenzzentrum mechanobiologie (atcz133) in the interreg v-a austria czech republic programme and by the institutional support of rvo: 68378297. this work was partially supported by the faculty of physics and applied computer science agh ust statutory tasks within subsidy of ministry of science and higher education. references [1] y. azizeh-mitra, h. m. enamul, p. r. g. s. v., u. nicholas. current strategies in multiphasic scaffold design for osteochondral tissue engineering: a review. journal of biomedical materials research part a 103(7):2460–2481, 2015. doi:10.1002/jbm.a.35356. [2] a. r. vaccaro, k. chiba, j. g. heller, et al. bone grafting alternatives in spinal surgery. the spine journal 2(3):206–215, 2002. doi:https://doi.org/10.1016/s1529-9430(02)00180-8. [3] p. koudelka, t. doktor, d. kytyr, et al. micromechanical properties of biocompatible materials for bone tissue engineering produced by direct 3d printing. key engineering materials 662:138–141, 2015. doi:10.4028/www.scientific.net/kem.662.138. [4] d. kytýř, t. doktor, o. jiroušek, et al. deformation behaviour of a natural-shaped bone scaffold. materiali in tehnologije 50(3):301–305, 2016. doi:10.17222/mit.2014.190. [5] i. kumpova, d. vavrik, t. fila, et al. high resolution micro-ct of low attenuating organic materials using large area photon-counting detector. journal of instrumentation 11(2), 2016. doi:10.1088/1748-0221/11/02/c02003. [6] a. nazarian, j. muller, d. zurakowski, et al. densitometric, morphometric and mechanical distributions in the human proximal femur. journal of biomechanics 40(11):2573–2579, 2007. doi:doi.org/10.1016/j.jbiomech.2006.11.022. [7] p. zlamal, o. jirousek, d. vavrik. a novel compression/tension device for investigation of trabecular bone failure using real-time micro-ct imaging. in 7th youth symposium on experimental solid mechanics, pp. 91–91. 2008. [8] v. rada, t. fila, p. zlamal, et al. multi-channel control system for in-situ laboratory loading devices. acta polytechnica ctu proceedings 18:15–19 2018. doi:10.14311/app.2018.18.0015. [9] t. doktor, d. kytyr, p. zlamal, et al. simulation of a three-point bending test on the isolated cell wall of aluminium foam. civil-comp proceedings 102, 2013. [10] bradski, gary. the opencv library, 2000. http://www.drdobbs.com/open-source/ the-opencv-library/184404319. [11] i. jandejsek, j. valach, d. vavrik. optimization and calibration of digital image correlation method. in proceedings of the 48th international scientific conference on experimental stress analysis, pp. 121–126. 2010. [12] b. d. lucas, t. kanade. iterative image registration technique with an application to stereo vision. vol. 2, pp. 674–679. 1981. [13] p. zlamal, o. jirousek, t. doktor, et al. compressive behaviour of trabecular tissue: finite element modelling and comparison using digital volume correlation. civil-comp proceedings 102, 2013. [14] t. fila, d. kytyr, p. zlamal, et al. high-resolution time-lapse tomography of rat vertebrae during compressive loading: deformation response analysis. journal of instrumentation 9(5), 2014. doi:10.1088/1748-0221/9/05/c05054. 27 http://dx.doi.org/10.1002/jbm.a.35356 http://dx.doi.org/https://doi.org/10.1016/s1529-9430(02)00180-8 http://dx.doi.org/10.4028/www.scientific.net/kem.662.138 http://dx.doi.org/10.17222/mit.2014.190 http://dx.doi.org/10.1088/1748-0221/11/02/c02003 http://dx.doi.org/doi.org/10.1016/j.jbiomech.2006.11.022 http://dx.doi.org/10.14311/app.2018.18.0015 http://www.drdobbs.com/open-source/the-opencv-library/184404319 http://www.drdobbs.com/open-source/the-opencv-library/184404319 http://dx.doi.org/10.1088/1748-0221/9/05/c05054 acta polytechnica ctu proceedings 18:24–27, 2018 1 introduction 2 materials and methods 2.1 sample preparation 2.2 tomographic scanning and reconstruction 2.3 additive manufacturing 2.4 uniaxial compression tests 2.5 strain measurement 3 results 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.8.0011 acta polytechnica ctu proceedings 8:11–13, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app dlc/ti thin films properties prepared by hybrid laser technologies jan mikšovskýa, b, ∗, miroslav jelíneka, b, petr písaříka, b, tomáš kocoureka, b, jan remsaa, c, karel jurekb a czech technical university in prague, faculty of biomedical engineering, nam. sitna 3105, kladno, czech republic b institute of physics ascr, na slovance 2, prague, czech republic c czech technical university in prague, faculty of nuclear sciences and physical engineering, břehová 7, prague, czech republic ∗ corresponding author: jan.miksovsky@fbmi.cvut.cz abstract. layers of diamond-like carbon are usable in many fields of industry as well as in medicine. many scientific groups have worked with different types of deposition techniques to prepare dlc layers with improved or unique properties. the dlc properties could be improved by various dopations. in this study, we focused on dlc layers doped by titanium, prepared by hybrid laser depositions. two techniques were used: dual pulse laser deposition (dualpld) and pulse laser deposition in combination with magnetron sputtering (pld/ms). preliminary tests for morphology, wettability, adhesion, hardness, corrosion, friction and wearability were examined. dlc samples were prepared on si(100) wafer and on ti6al4v alloy substrates with titanium concentration from pure up to 25 at.%. friction of the prepared layers ranged from 0.09 to 0.18. the films exhibited very low wear for loads 1 n and 2 n. the tested pin wear grew with decreasing titanium dopation. the measurements of the contact angle and determination of surface free energy were done for water, diiodomethane and ethylene glycol. hardness and young's modulus decreased with higher titanium dopation, on the other hand adhesion improved with increasing amount of titanium in film. keywords: pld, dlc, titanium, surface properties, laser deposition, magnetron sputtering. 1. introduction in this study, we focused on the changes in the mechanical and corrosion behavior with various amount of titanium in dlc/ti thin films. two techniques for dopation of titanium were used [1–3]. the first technique was based on pulsed laser deposition (excimer laser, krf, λ = 248 nm) combined with dc magnetron sputtering. the second technique was dualpld. laser fluency was 8 j/cm2 for dlc deposition and 5 j/cm2 for metal dopant. power of magnetron was from 40 w to 300 w for metal dopant. we studied behavior of layers for ti concentration from 1 at.% to 25 at.%. si(100) and ti6al4v alloy were used as substrates. prepared films were characterized with focus on their basic physical properties, such as crystallinity by xrd, morphology by afm and mechanical profilometry, chemical bonding nature by wdx, surface free energy by contact angle. hardness and young’s modulus were studied by nanoindentation. macro adhesion was studied up to 30 n. wearability and tribological properties were studied as well. the content of sp3 bonding was studied by xps. special focus was given on tribological properties. corrosion rates were measured as function of ti concentration. 2. experimental wearability test was performed by tribometer (anton paar) with linear movement. chromium steel testing ball with 6 mm diameter was used. speed of the testing ball was 5 cm/s, total testing length 10 m, used loads during tests were 1 n and 2 n. the tests were performed as dry. the friction was determined from the test records. nanoindentation tests were performed on all samples for determination of hardness and elastic modulus. all measurements were implemented on layers with thickness 500 nm and the indentation depth was less than 10 % of the layers thickness. for more details, see [4]. adhesion of the films was tested for samples on ti6al4v substrates. we tested samples by progressive load from 1 n to 30 n. three types of critical forces were evaluated. at least two scratches for each layer were tested. for more details, see [4]. wettability was measured by three liquids: distilled water, diiodomethane and ethylene glycol, by contact angle measurement system (dsa100, krüss company). surface free energy was determined by drop shape analysis software with fowkes method. the volume of droplets was (0.4 ± 0.1) ml. the measurement technique was sessile droplet and static contact angle. each sample was tested at least on three 11 http://dx.doi.org/10.14311/app.2017.8.0011 http://ojs.cvut.cz/ojs/index.php/app j. mikšovský, m. jelínek, p. písařík et al. acta polytechnica ctu proceedings figure 1. from left wearability of the ti6al4v substrate and samples ti:dlc7(pure dlc), ti:dlc10 (5 at.% ti) and ti:dlc12(25 at.% ti). the figures are captured at magnification 200x. figure 2. wearability of the testing ball after tests for ti6al4v substrate and for samples ti:dlc7(pure dlc), ti:dlc10 (5 at.% ti) and ti:dlc12(25 at.% ti). the figures are captured at magnification 200x. dual pld 0 at.% ti 1 at.% ti 3 at.% ti 5 at.% ti 10 at.% ti 25 at.% ti 0.12 0.14 0.13 0.12 0.11 0.09 pld/ms 0 at.% ti 1 at.% ti 3 at.% ti 5 at.% ti 10 at.% ti 25 at.% ti 0.13 0.15 0.14 0.18 0.15 0.16 table 1. friction on the layers of ti:dlc prepared by dualpld and pld/ms methods. different places. for each liquid, a new sample was used. from each droplet, series of 15 pictures were evaluated. laplace profile fitting was used. tested samples were on si (100) and thickness of the films was around 100 nm. we finished preliminary tests of the corrosion behavior as function of the amount of the titanium dopant in the dlc layers. we used potentiodynamic measurement and tafel plots for icorr determination and then for corrosion rate determination. 3. results contents of the titanium dopant in the dlc layers were determined by wdx and confirmed by xps. prepared samples concentrations were 0, 1, 3, 5, 10 and 25 at.% of titanium. dualpld technique was better to reach the desired dopants concentration. the xrd tests confirmed that all the samples were amorphous with detection of ti and tic phases for the highest ti concentration (10 and 25 at.% ti). thickness of the prepared layers were approximately 500 nm, for layers on ti6al4v substrates and 100 nm for layers on si (100). afm was used for morphology determination. both methods produced smooth layers around 1 nm, just for dualpld method roughness increased for 10 and 25 at.% ti up to 5 resp. 15 nm. wearability tests have been done for loads 1 n and 2 n. for these loads, there were no damage observed on samples. the wearability of the ball decreased with increased amount of titanium in layers. examples are on the figures 1 and 2. from comparison of the coated and uncoated materials, the uncoated materials were seriously damaged. the friction for pure dlc layers was 0.125 ± 0.05. for dual pld method we observed slight increase in friction for lower concentration up to 0.14 and then decrease to 0.09. for pld/ms method we observed slight increase for dopated layers but without clear trend, for more detail see table 1. the hardness for pure dlc layers starts at (31 ± 1) gpa and was decreasing with increasing of the titanium content to 12 ± 2 gpa for 25 at.% of titanium. the decrease was slightly quicker for pld/ms technique. the same behavior we observed for elastic modulus with the highest values around 250 gpa for pure dlc and decrease to 150 gpa for 25 at.% ti content of titanium practically same for both methods. the adhesion values seemed to increase with content of the titanium with exception for highest concentration of titanium for pld/ms method. critical force fc3 started at 5.6 n for pure dlc and increased to 18.4 n for dualpld technique for 25 at.% ti. for pld/ms technique critical force was increasing up to 20.2 n for 10 at.% ti but for 25 at.% ti it reduced to 9.2 n. this could be connected with lower hardness of titanium. the contact angle for water was in range from 57.7° to 71.5° for dual pld method and from 53.8° to 70.0° for pld/ms method. ca for ethylene glycol was in range from 34.5° to 42.8° resp. from 30.1° to 44.0° and for diiodomethane the range was from 12 vol. 8/2017 dlc/ti thin films properties prepared by hybrid laser technologies 34.9° to 38.3° resp. 32.0° to 36.4° for dual pld resp. pld/ms method. the results showed no trends in ca in correlation with titanium content in layers. from the free surface energy calculation we can conclude that all the tested layers are dominantly disperse. from the corrosion tests, the high concentration of the titanium dopants seemed to improve corrosion resistivity against the pure dlc. for various concentrations there were no clear trends observed and further tests should be performed before any conclusions could be made. 4. conclusion we prepared diamond-like carbon samples doped by titanium in concentration up to 25 at.% by two deposition techniques dual pld and pld/ms. we made standard testing as xrd, wdx, xps, afm, profilometry measurements etc. the main focus was put on mechanical and tribology testing. for both techniques, the hardness and elastic modulus were decreasing with higher content of the titanium, as we expected. there were no significant changes based on used methods. we observed better adhesion related to increasing titanium concentration in films. this is connected with lower compression stress due dopation. on the other hand, for the highest concentration for pld/ms there was lower adhesion observed, which could be connected with lower hardness. wearability test showed no damage to the coatings at load 1 n or 2 n (no matter on ti concentration) but wear of the testing ball was significantly higher for lower ti concentration in films. this is probably connected to higher layer hardness. tests will be continued with higher loads. for corrosion behavior, only preliminary tests are available. corrosion resistance was better for higher titanium amounts, when the highest concentration had up to three times better icorr parameter. for wettability, there are no clear trends based on titanium content. all layers are dominantly disperse. acknowledgements this work was supported by the grant agency of the czech technical university in prague, sgs16/110/ohk4/1t/17, sgs16/190/ohk4/2t/17, by meys cr grant lo1409 and by czech grant agency ga15-05864s. references [1] j. bulíř, m. novotný, m. jelínek, et al. plasma study and deposition of dlc/tic/ti multilayer structures using technique combining pulsed laser deposition and magnetron sputtering. surface and coatings technology 200(1 4):708 – 711, 2005. doi:http://dx.doi.org/10.1016/j.surfcoat.2005.01.087. [2] q. wei, r. narayan, j. narayan, et al. improvement of wear resistance of pulsed laser deposited diamond-like carbon films through incorporation of metals. materials science and engineering: b 53(3):262 – 266, 1998. doi:http://dx.doi.org/10.1016/s0921-5107(98)00150-0. [3] a. voevodin, m. capano, s. laube, et al. design of a ti/tic/dlc functionally gradient coating based on studies of structural transitions in ti-c thin films. thin solid films 298(1 2):107 – 115, 1997. doi:http://dx.doi.org/10.1016/s0040-6090(96)09145-6. [4] m. jelínek, j. zemek, t. kocourek, et al. dual laser deposition of ti:dlc composite for implants. laser physics 26(10):105605, 2016. 13 http://dx.doi.org/http://dx.doi.org/10.1016/j.surfcoat.2005.01.087 http://dx.doi.org/http://dx.doi.org/10.1016/s0921-5107(98)00150-0 http://dx.doi.org/http://dx.doi.org/10.1016/s0040-6090(96)09145-6 acta polytechnica ctu proceedings 8:11–13, 2017 1 introduction 2 experimental 3 results 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0069 acta polytechnica ctu proceedings 7:69–71, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app performance of steel fiber reinforced concrete – comparability of tests according to dafstb-guideline "stahlfaserbeton" and en 14651 steffen anders∗, melanie schovenberg institute of structural engineering, bergische universität wuppertal, pauluskirchstr. 11, 42285, wuppertal, germany ∗ corresponding author: baustoff@uni-wuppertal.de abstract. for the determination of the performance of steel-fiber reinforced concrete (sfrc), the post-peak flexural strengths are used. in different national and european standards, different test-setups are defined, resulting in double efforts for testing for the manufacturers. in addition, the german national guideline "stahlfaserbeton (dafstb)" on sfrc is well established european-wide, but the test standard is specifically national, demanding a four-point-bending tests using unnotched beams. contrarily, the european standard en 14651 as well as the model code 2010 [1] demand three-point bending tests using notched specimens. applying the national guideline is obligatory in germany for structural use of sfrc. therefore, it is essential to standardize the performance evaluation of sfrc based on commonly applied international guidelines. in the following, an approach is presented especially dealing with the problem of random occurrence of cracks in the four-point-bending tests. it is shown, that neglecting the point of crack can systematically under-estimate the performance of sfrc especially at deformations. keywords: steel-fiber reinforced concrete, performance classes, post-peak flexural strength. 1. introduction in order to get permission to distribute and sell steelfibers in europe, a ce-certificate is obligatory. therefore, it is necessary to determine the performance of the steel-fibers in concrete, depending on the type of steel fibers as well as the fiber contents. in this case, the en 14651 [2] is the relevant standard, demanding a test using third-point loading on notched specimens. in order to get the permission to sell steel-fiber reinforced concrete e.g. as ready mixed concrete the performance of sfrc has to be evaluated according to the german national guideline "stahlfaserbeton (dafstb)" [3]. this guideline demands tests on unnotched specimens in a 4-point bending set-up. in both guidelines, performance of sfrc is measured in terms of post-peak flexural strengths. seen from the mechanical point of view the evaluation of post-peak flexural strengths is a fracture mechanical problem, demanding determination of strengths and load bearing capacities directly in the crack. this precondition is fulfilled only by en 14651 but not by the german guideline, because the crack can occur anywhere in the mid third of the beam due to the four-point-loading scheme. in order to overcome the problem of mechanical incorrectness as well as achieving international standardization, establishing a connection between the two test methods would be advantageous to facilitate the conversion of results from one method to another. one important aspect in this context is to include the point of crack into the evaluation according to the guideline "stahlfaserbeton". this aspect will be dealt with in the following. 2. test methods and evaluation of post-peak flexural strengths the most important difference between the national guideline "stahlfaserbeton (dafstb)" and the european standard en 14651 is the test set-up as shown in fig. 1. in the four-point loading test, which is demanded by the guideline "stahlfaserbeton (dafstb)", the crack should occur randomly in the mid third between the two loads due to lacking of a notch. the deflection of the beam is measured using transducers at mid-span. as described, due to this distance the testing method is not consistent from a fracture mechanical point of view. the method given in en 14651 uses a notched beam and just one load exactly above the notch at mid-span. thus, deflection and cracking are in-line. the evaluation of the post-peak flexural strength in the guideline "stahlfaserbeton (dafstb)" uses the mid-span deflection δ. in this guideline, the post-peak flexural strength is calculated from the load at two defined deflections of δ = 0.5 mm and δ = 3.5 mm at mid-span (see fig. 2-up). seen from point of view of fracture mechanics, the deflection at the point of crack is always higher than the deflection at mid-span, resulting in a systematic underestimation of the postpeak flexural strength, provided a strain softening behavior of the steel-fiber reinforced concrete. 69 http://dx.doi.org/10.14311/app.2017.7.0069 http://ojs.cvut.cz/ojs/index.php/app steffen anders, melanie schovenberg acta polytechnica ctu proceedings figure 1. comparison of test set-ups on the other hand, the en 14651 standard uses the crack mouth opening displacement (cmod) being directly correlated to the mid-span deflection [4]. to specify the post-peak flexural behavior, strength values at four different points are defined in terms of cmod (see fig. 2-down). the deflection at δ = 0.47 mm at cmod1 correlates well to the german guideline. cmod4 at a deflection of δ = 3.02 mm shows a difference of about 0.5 mm for the evaluation point. additionally, the cross-sections differs either, for the notch reduces the height of the cross section (both beams are produced with a cross-section of 150 × 150 mm). the approach of the en 14651 is consistent with respect to fracture mechanics, because the deflection is measured directly in the cracked crosssection. the shown differences in testing and calculating make it nearly impossible to compare the results of the guidelines. a special problem being the random cracking with german test set-up. 3. calculation approach to adjust the results taken during the tests according to the guideline "stahlfaserbeton (dafstb)" the point of the first crack has to be measured. if this is done the deformations measured at mid-span, can be transferred to the cracked cross-section by simply using theorem of intersecting lines applied on the larger part of the beam, assuming both parts of beam are not deformed themselves. fig. 3 shows that the deflection in the cracked cross-section is always larger than deflection at mid-span. the consideration of the point of the crack results in an increased stretching of the force-deflection curve. the correction of the curve increases with an increasing deflection during the test as well as with an infigure 2. evaluation of post-peak flexural strengths at different points. up dafstb-guideline "stahlfaserbeton", down en 14651 figure 3. calculation approach for taking the point of crack into account using the theorem of intersecting lines creasing distance of the crack from the middle of the beam. in fig. 4, the increased stretching can be easily seen especially at higher deflections. figure 4. change of mid-span deflection in case of strain hardening material at low deflections and strain softening at high deflections 70 vol. 7/2017 performance of steel fiber reinforced concrete depending on the effectiveness of the fibers (and on the fiber content), generally two different material behaviors are distinguished. either the material is strain hardening, meaning that the transferred load increases with an increasing deflection or the material is strain softening, showing decreasing transferable loads with increasing deflections. in fig. 4 the postpeak flexural strength for low deflections of about δ = 0.5 mm shows strain hardening. in case of strain softening, as can be seen at a deflection higher than about 1 mm a correction of the load-deflection curve leads to increasing post-peak flexural strengths. 4. results and discussion after measuring the point of crack of about 360 fourpoint bending tests, the resulting distribution of the point of crack is given in fig. 5. the displayed distance of the crack from mid-span does not differ whether the crack occurred left or right from the mid of the beam. assuming a perfectly homogeneous concrete, the crack should theoretically occur between the loads. an analysis of the point of the first crack shows, that only 50 % of all cracks occur in a distance less than 4 cm around the mid of the beam. at distances of less than 2 cm from mid-span, less than 25 % of all cracks occur. figure 5. distance of point of crack from mid-span applying the correction approach for the point of crack on several beams, the post-peak flexural strength values are affected as well. the larger the distance of the crack from mid-span the higher the correction, strain softening assumed. most of the beams have shown force-deflection curves comparable to fig. 4. as mentioned, these beams often show little strain hardening at low deflections. because the correction at δ = 0.5 mm is small, the post-peak flexural strength values are unaffected for a large number of beams. due to the slight strain hardening, the post-peak flexural strength values decrease for a number of beams. at larger deflections of δ = 3.5 mm the correction is higher, because of the higher deflection in combination with the strain softening. the mean increase ranges at about 0.15 n · mm−2, which is half a performance class. figure 6. effect of correction of post-peak flexural strength due to random cracking of beams tested according to dafstb guideline "stahlfaserbeton" 5. conclusions for the determination of performance of sfrc, different testing and evaluation approaches are used in germany and in europe. from a point of view of fracture mechanics, the german approach is incorrect, due to a random occurrence of the cracks predominantly in the mid-third. this leads to a systematic underestimation of the deflection in the cracked cross-section. therefore, it is essential to take the point of the first crack into account, leading to a stretching of the forcedeflection curve. two different modes of behavior can occur. for strain hardening, which is often found at low deflections and at higher fiber contents, the postpeak flexural strength can decrease, and for strain softening behavior, usually found for concrete having lower fiber contents as well as at higher deflections, the post peak flexural strength increases. taking the point of crack into account leads to an increase of about half a performance class on an average. in further steps, the results should directly be related to results from tests according to en 14651. additionally, the different geometries caused by the notches have to be taken into account. references [1] din en 14651: prüfverfahren für beton mit metallischen fasern bestimmung der biegezuggfestigkeit (proportionalitätsgrenze, residuelle biegezugfestigkeit) 2004. [2] deutscher ausschuss für stahlbeton (dafstb): dafstb-richtlinie stahlfaserbeton 2012. [3] fédération international du béton (fib): modelcode for concrete structures 2010. [4] b. barr, et al. round-robin analysis of the rilem tc 162-tdf beam-bending test: part 2 application of delta from the cmod response. materials and structures 36(263):621–630, 2003. doi:10.1617/13954. 71 http://dx.doi.org/10.1617/13954 acta polytechnica ctu proceedings 7:69–71, 2017 1 introduction 2 test methods and evaluation of post-peak flexural strengths 3 calculation approach 4 results and discussion 5 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0104 acta polytechnica ctu proceedings 12:104–107, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app the future of passenger air transport – very large aircraft and out key human factors affecting the operation and safety of passenger air transport petra skolilová ctu in prague, faculty of transportation sciences, czech republic correspondence: skolilova@fd.cvut.cz abstract. the article outlines some human factors affecting the operation and safety of passenger air transport given the massive increase in the use of the vla. decrease of the impact of the co2 world emissions is one of the key goals for the new aircraft design. the main wave is going to reduce the burned fuel. therefore, the eco-efficiency engines combined with reasonable economic operation of the aircraft are very important from an aviation perspective. the prediction for the year 2030 says that about 90% of people, which will use long-haul flights to fly between big cities. so, the a380 was designed exactly for this time period, with a focus on the right capacity, right operating cost and right fuel burn per seat. there is no aircraft today with better fuel burn combined with eco-efficiency per seat, than the a380. the very large aircrafts (vlas) are the future of the commercial passenger aviation. operating cost versus safety or co2 emissions versus increasing automation inside the new generation aircraft. almost 80% of the world aircraft accidents are caused by human error based on wrong action, reaction or final decision of pilots, the catastrophic failures of aircraft systems, or air traffic control errors are not so frequent. so, we are at the beginning of a new age in passenger aviation and the role of the human factor is more important than ever. keywords: vla, human factors, air transport, air traffic, security. 1. introduction the study of human factors deals with people’s interaction with their environments. in the case of general aviation, this is the study of pilot performance with influence on the emotions, stress, fatigue or design of cockpits or interaction with crewmembers, air traffic control personnel and others. fatigue is an expected aspect of life and there is a very thin border between normal daily fatigue (for example fatigue from lack of sleep) and the tough fatigue from continuous stress. typically, there are no significant marks or consequences. but the aftermath in general aviation environment can be disastrous[1]. 1.1. introduction of the air traffic markets situation aviation mega-cities (amcs) are the new phenomenon for the long-haul market. for long-haul flights we are rating the traffic over 2,000 nautical miles. more than 40% of all long-haul passengers are beginning or ending, at amcs and another 20% are between these 40 amcs. more than another 40 cities have a potential to grow to amcs. by the gmf of airbus, the a380 is the choice number one for these markets. the current markets situation is represented figure 1. the wide-body aircraft is a very important aspect of the aircraft market as well. these aircraft types are more focused on the routes for amcs than to the secondary cities. this is the right market opportunity for the highly efficient new a350xwb as well as the a330neo by the airbus. 1.2. very large aircraft according to airbus statistics before the vlas age, one million passengers per month were served by 200 aircrafts per day. but today, it is only 100 vla aircrafts connecting the world’s major hub. the passengers prefer to fly on a vla due to the high travel comfort which only the vla aircraft can provide. one of the most comfortable seating configurations is the 407 pax (passengers) on board configuration. but also, the maximum capacity of 853 seats are a highly desired option for airlines, which means they have the flexibility to mix comfort with revenue and cost per seat. but today, a typical configuration of an a380 is between 500-520 seats in three or four traveling class configurations. the capacity flexibility is one of the best airline tools how to optimize the right capacity at the right time between mature and expanding markets and their reaction you can read from figure 2[2]. we can find three key reasons why the vlas aircraft are the future of aviation: a) solution to sky congestion: reducing the number of aircrafts needed for transport of the ever-growing air passenger volume; b) solution to the changing demand at different times of the day; c) solution to overcrowded slots at the megacities airports. the standard solution, which is often implemented 104 http://dx.doi.org/10.14311/app.2017.12.0104 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 the future of passenger air transport – very large aircraft figure 1. concentration of flights in long-haul markets [source: oag, airbus]. figure 2. new deliveries of vla by region [source: oag, airbus]. 105 petra skolilová acta polytechnica ctu proceedings on routes via the a380, is operating the a380 during peak hours and other aircraft types at non-peak hours during the day on the same routes. in fact, 50% of airports-pairs served by the a380 have another aircraft type in operation also. for the airlines it means to target the right capacity at the right time during the day, for example a combination of the a380 and a350xwb aircraft. the main markets for the vla are the long-haul markets between amcs. in these days they are mainly on the asia/pacific area, where more than 25% flights are served by the a380[2–4]. "from today until 2033, we forecast the demand for nearly 1,230 vla globally. the two largest markets are asia-pacific and the middle east. asia-pacific will represent 47% of the demand for vla over the next 20 years which can be clearly seen by the large number of vla today flying to from and within asiapacific followed by the middle east which today has the largest fleet and largest backlog"[2]. by the gmf prediction we can await an annual growth of the passenger air traffic around 4.7% and it will assist approximately 30,555 passengers’ aircraft during next 20 years. "this demand represents $4.4 trillion at book value over the next 20 years. of this demand for passenger aircraft, 40% will be for the replacement of ageing, less fuel efficient aircraft and 60% of the demand will be for growth"[3]. 2. human factors training in aviation safety means different things to different people. to the traveling public, the term ’safety’ means, at its most fundamental level, that we want to reach our destinations without getting hurt. since people do get hurt on occasion, it is logical to conclude that we are willing to accept some risk in traveling. whether consciously or subconsciously, we know that there is some chance, albeit a minute one, that we could be hurt or killed while traveling. today we average about 50 accidents in commercial airlines each year, but it means that the odds of any one flight being your last one are remote ( 1 to 4,9 million ) and given that more people are killed or injured in automobile accidents (approximately 44,000 deaths per year), ’safe’ is a reasonable word to apply to commercial aviation[5]. one of the first airlines to accept the human factors training was continental airlines. they accepted the existing format and settled the special training called cockpit resource management, evolved into their original training initially titled "crew coordination concepts", and later changed to "maintenance resource management"[6]. it covers communications with pilots, pilots with cabin crewmembers, labor with management, and companies with the government. the icao (international civil aviation organization) is a specialized agency of the united nations, which codifies the principles and techniques that aviation has developed and implemented a special human factors training called sms program: safety management system. the definition of sms: "a dynamic risk management system based on quality management system (qms) principles in a structure scaled appropriately to the operational risk, applied in a safety culture environment." 2.1. human reliability analysis (hra) it is well-established that the largest contribution to the probability of system failure in complex systems is human error. a risk assessment method of any value in aviation must consider human interactions and errors on system reliability. hra models human behavior to better understand human interactions with the rest of the system. a popular method used for human performance reliability prediction is the technique for human error rate prediction (therp). therp are model with five steps: (1.) define the system or process. (2.) identify and enumerate all human operations performed and their relationships to the system. (3.) predict error rates for each human operation. (4.) determine the effect of human error on the system. (5.) make changes that will reduce the system failure rate. pra has been used in chemical, transport, energy, aerospace and military. the first step in an hra is a study to identify the hazards and their impact on the scale. the scale is uses fault trees (ft). the hazards are assigned probabilities, which are propagated up the tree to establish the probability of the undesired top hazards. in practice, it means that the phase of mapping the hazards and human failures is followed by a practical modeling scenario with deductive logic and probabilistic tools called fault trees (ft). the ft are composed of three levels: the top level called type of accident which defines the accident scenario, the middle level are the intermediate failures causing the top level and the basic level are the failures that cause the top level to occur. developing the scenario is a highly sophisticated method with extreme dependency on source data of the model and good subject knowledge. it could be supported by probabilistic software tools. 3. conclusion in these days, more than 300 airbus a-380 are in service operating under the flags of 13 carriers: singapore airlines, emirates, qantas, air france, lufthansa, korean air, china southern airlines, malaysia airlines, thai airways international, british airways, asiana airlines, qatar airways and etihad airways using airports all over the world. one part of responsibility for passengers’ security is on airport management, how to manage the airports runways, taxiways, gate areas 106 vol. 12/2017 the future of passenger air transport – very large aircraft and terminals to accommodate this large and heavy aircraft and more than 700 pax per flight. what has resulted has been a relatively seamless, trouble-free adaptation on the part of airports, which reflects the results of thoughtful planning and analysis. the other part is the human factor. commercial air carriers continue to restructure. the world’s economic downturn beginning in 2008 forced many air carriers out of business and other to consolidate. as a result, air traffic management has had to, and will continue to, become more flexible in its contractual obligations with air carriers, their aeronautical revenue structures, and increasingly the use of their facilities. in the past, aviation safety improvements were characterized by a fly-crash-fix-fly approach. sometimes the causes would be weather-related or a mechanical failure, but more often the causes would be determined to be human factor error – usually the pilot. today we realize that it is much more productive to engineer a system in which causes of failure have been prevented. as one might imagine, there are many elements to this engineering effort. the understanding of hazard identification, risk management, system theory, human factors engineering, organizational culture, quality engineering and management, quantitative methods and decision theory must have a pillar in aviation development. the requirement for vlas will grow with the world’s network and as more people have the ability to fly. by 2032, this demand will result in a need for more than 1,300 vlas. given the projected growth in asia-pacific, both economic and air passenger traffic growth, the region’s demographics and urbanization trends, it is unsurprising that the region’s airlines will take 47% of these aircraft over the next 20 years. the middle east will be the second largest region in terms of demand for vlas, at 26%. this can be seen today in the size of the backlog of a380s within middle eastern carriers[3]. references [1] s. young. airport: planning & management 6th edition. mcgraw-hill companies, inc., 2011. [2] a. 2015. airbus. [online], přístup získán may 2015. [3] iata. the international air transport association. [online]. available: http://www.iata.org/about/pages/index.aspx. accessed mar 2015. [4] wttc. world travel & tourism council. [online]. available: www.wttc.org. přístup získán mar 2015. [5] the economist group. [online]. available: http://www.eiu.com/home.aspx. přístup získán apr 2015. [6] r. stanley, r. trollip. human factors for general aviation. jeppesen sanderson, 2011. 107 http://www.iata.org/about/pages/index.aspx www.wttc.org http://www.eiu.com/home.aspx acta polytechnica ctu proceedings 12:104–107, 2017 1 introduction 1.1 introduction of the air traffic markets situation 1.2 very large aircraft 2 human factors training in aviation 2.1 human reliability analysis (hra) 3 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0011 acta polytechnica ctu proceedings 13:11–15, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app performance of cement composites reinforced with surface-modified polypropylene microand macro-fibers jakub antoša, ∗, lukáš dejdara, jan trejbala, zdeněk prošeka, b a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b czech academy of science, institute of physics, na slovance 1999/2, 182 21 prague 8, czech republic ∗ corresponding author: jakub.antos@fsv.cvut.cz abstract. this paper focuses on the mechanical properties investigation of cement pastes reinforced with surface treated polymer fibers. the cement matrix was composed of portland cement (cem i 42.5 r, w/c ratio equal to 0.4). two polypropylene fiber types (microand macro-fibers) were used as randomly distributed and oriented reinforcement in volume amount of 2 %. the fibers were modified in the low-pressure inductively coupled cold oxygen plasma in order to enhance their surface interaction with the cement matrix. the investigated composite mechanical properties (load bearing capacity and response during loading) were examined indirectly by means of four-point bending mechanical destructive tests. a response of loaded samples containing treated fibers were compared to samples with reference fibers. moreover, cracking behavior development was monitored using digital image correlation (dic). this method enabled to record the micro-cracks system evaluation of both fiber reinforced samples. keywords: cement composite, fiber reinforcement, polymeric fiber, surface modification, plasma treatment, four-point bending, dic, crack pattern. 1. introduction plain concrete can be described as a tension-weak material, where cracks often occur due to exceed of maximal tensile stress, rapid drying, shrinkage, and the like. the crack system development generally increases in the time and thus gradually propagates the concrete volume. as moisture (containing substances and ions), co2 and other degrading agents comes through formed cracks, service-life, mechanical performance and other properties of the concrete degrade [1]. to reduce all above mentioned phenomena, the utilization of fibers seems to be an effective possibility to control creation, development and width of cracks [2]. during the mixing process, fibers are randomly oriented and distributed throughout the concrete in all directions. consequently, the fresh concrete is thus protected against excessive formation of plastic shrinkage cracks (concrete tensile strength as well as modulus of elasticity thus increased). in the hardened concrete, the fibers prevent evolution of micro-crack system into macro-crack formation. this is ensured by redistribution of tension stress through the fibers “bridging” a crack (sometimes refereed to as “sewing” effect) and preventing disintegration of broken concrete [1]. in the civil engineering practice, there are several types of fibers used as scattered reinforcements in concrete, e.g. metal, polymer, glass or made of various other materials. the polymer fibers (polypropylene or polypropylene-based) are often preferred over steel fibers due to lower production cost, increased resistance to degradation and sufficient mechanical properties (such as tensile strength or toughness) [3]. moreover, polymeric fibers provide an increase of fire resistance for the structure due to low melting point of the polymers and inferential formation of water vapor capillaries in the structure of concrete [4, 5]. on the other hand, due to low chemical activity of most polymers (related to cement paste), the chemical interaction (bonding) between fiber surfaces and cement matrix is not ensured. the hydrophobic character of fiber surface does not create strong bond with polar liquid (water contained in cement paste). the inter-facial transition zone is thus quite porous, as captured in fig. 1. moreover, smooth surfaces of fiber provides insufficient mechanical interaction between the materials. in general, strong inter-phase interaction between matrix and reinforcement is crucial to ensure mechanical (and other) properties of composites [6]. in order to enhance the physico-chemical interaction between the two materials, several types of fiber surface modifications including chemical or physical treatment can be applied [7]. in most cases, the standard treatments have strong impact on fibers mechanical properties. moreover, the chemical treatment is not environmentally friendly and is too time and energy consuming (an exposition fiber to alkali solution lasts up to tens of hours and whole process requires high temperature) [8]. 11 http://dx.doi.org/10.14311/app.2017.13.0011 http://ojs.cvut.cz/ojs/index.php/app j. antoš, l. dejdar, j. trejbal, z. prošek acta polytechnica ctu proceedings figure 1. porous inter-facial transition zone between mono-filament fiber and cement matrix [7]. the plasma treatment seems to be an effective and thrifty method to modify both physical and chemical polymer surface. the physical treatment rests in a roughening via ion bombardment. unlike the conventional physical methods, this influences only the monomolecular layer, so fiber mechanical performance remains almost untouched. next to that, the chemical modifications rest in an activation of surfaces due to interaction between polymeric chains and free radicals. the active polar groups are implemented onto fiber surfaces and thus change hydrophobic characteristic to hydrophilic [9, 10]. the most suitable time to modify polymeric micro-fiber using plasma treatment is 30 seconds according to [11, 12]. in presented study, we focused on inter-face interaction between reference and plasma treated fibers with cement matrix. we modified two fiber types – polypropylene microand macro-fibers, as the common representatives of reinforcement in the field of concrete constructions. we modified the fibers by low-pressure cool oxygen plasma treatment. as an indirect indicator of interaction between the two materials, we have done four-point bending tests of fiber reinforced prismatic cement specimens when we focused on their mechanical response. moreover, we recorded the specimens behavior during loading using digital image correlation. 2. materials 2.1. polymeric fibers two types of polypropylene (pp) fibers differing in dimensions were used. micro-fibers having 18 µm in diameter and approx. 12 mm of length (commercial name: stavon, manufacturer: trevos, s. r. o., czech republic) were used for reinforcement of the first set of cement composite samples. these fibers were surface treated by a fabric softener. according to product data sheet, material properties of the fibers are declared as: uts ≈ 273 mpa (ultimate tensile strength calculated from dtex unit) and ductility = figure 2. top – polypropylene macro-fibers, bottom – polypropylene micro-fibers. min. 50 %. no other information about the fibers is provided by the manufacturer. for the second set of experiment samples, the dimensions of used fibers were 80 µm in diameter and ≈ 19 mm of length (commercial name: highgrade, manufacturer: brugg contec ag, switzerland). the basic material properties listed in data sheet of the product are as follows: % = 910 kg m−3 (volumetric mass density), uts = 400 mpa, e = 4.9 gpa (young’s modulus of elasticity), mp ≈ 150 °c (melting point). according to european standard code [5], cementitious composite reinforced by above described macro-fibers is classified as construction concrete, not plain. both types of fibers are captured in figure 2. 2.2. cement paste prismatic specimens composed of cement paste (cem i 42.5 r, water to cement ratio 0.4) and polymer fiber reinforcement were made using metal forms of 40×40×160 mm dimensions. on the whole, four different mixtures were prepared (reference and treated, microand macro-fibers). the amount of fiber reinforcement was equal to 2 % of sample volume. after 1 day, the prepared samples were cast out from forms and stored in a water environment to prevent degradation for other 27 days (total age of samples equal to standard 28 days), when the samples were tested. 12 vol. 13/2017 polymeric fiber reinforced cement composite figure 3. crack development during four-point bending test of cement sample reinforced with reference (top) and 30 seconds plasma treated (bottom) microfibers. figure 4. crack development during four-point bending test of cement sample reinforced with reference (top) and 30 seconds plasma treated (bottom) macrofibers. 3. experimental methods 3.1. plasma treatment both macroand micro-fibers were modified in order to change their surface properties by means of plasma treatment. concretely, low-pressure cool oxygen coupled plasma treatment was done using tesla vt 214 (13.56 mhz) device. the plasma treatment duration was equal to 30 s. other treatment parameters were as follows: 100 w total power of rf source, 20 pa total gas pressure, 50 sccm o2 flow. after plasma treatment, the fibers were admixed into fresh cement paste immediately. 3.2. four-point bending test the interface interaction between fibers and cement matrix was assessed by the destructive four-point bending tests. the indirect testing was executed as reimbursement of standard single-fiber pull-out test due to complicated fiber dimensions and shapes. each type of mixture was represented by 6 specimens in order to obtain statistically relevant set of results. the prismatic cementitious composite specimens were tested on loading frame web tiv ravestein fp 100. base-span was equal to 120 mm. the test was displacement controlled with 1 mm/min rate of loading and it was interrupted when the mid-span displacement reached 4 mm and 3 mm in the case of samples reinforced with microand macro-fibers, respectively. the main focus of the experiment was on the post cracking phase with expected increase of sample toughness. 3.3. digital image correlation the behavior of cracks was observed using digital image analysis (dic). dic analysis allowed to image the amount, shape and position of cracks graphically and thus answer the question how the specimens resisted to loading in the post-cracking phase (single vs. multiple cracking behavior) [14]. as the result of the analysis, cracking patterns were visualized with a map of maximum principal strain in which concentrations can be attributed the crack formation. to allow the dic observation, the samples were equipped with artificial stochastic contrast pattern onto their surfaces. images used for dic analysis were taken with dslr camera canon eos 70d fitted with canon zoom lens ef-s 17-85 is usm. the exposition parameters were as follows, shooting interval 10 seconds, iso speed 100, focal length 50 mm, aperture value 4.5, shutter speed 1/100 sec. 4. results 4.1. samples reinforced with micro-fibers it was shown, that maximum load-bearing capacity of cement samples reinforced with micro-fibers (d = 18 µm) was lower than 7 kn when reference fibers were used and almost 8 kn in the case of plasma modified fibers. the residual toughness of specimens containing surface treated fibers in the post cracking phase (namely after 3 mm mid-span displacement) was significantly higher, if compared to the referential specimens. this phenomenon can be attributed to increase of adhesion between pulling-out fibers from matrix when sample underwent deformation. nevertheless, slight increase of initial toughness in the linear-elastic phase of loaded samples was also recorded. it was probably caused by reduction of amount and wight of shrinkage cracks. the force-displacement diagram is shown in fig. 5. dic analysis showed that loaded cement sample was damaged by one single crack, which was opened when the displacement was increased in the case of reference fibers. on the other hand, sample reinforced with surface treated fibers was damaged by systems of cracks, known as multiple cracking behavior. the principal strain diagrams are shown in the fig. 3. 13 j. antoš, l. dejdar, j. trejbal, z. prošek acta polytechnica ctu proceedings figure 5. force-displacement diagram of cement samples containing micro-fibers. figure 6. force-displacement diagram of cement samples containing micro-fibers. 4.2. samples reinforced with macro-fibers loading response of cement samples reinforced with macro-fibers (d = 80 µm) exhibited similar results compared to micro-fiber reinforcement. the maximum reached force of reference specimens can be found close to 6 kn. the plasma treated fiber-reinforced samples reached load-bearing capacity close to 8 kn. however, the residual toughness of samples beyond mid-span deflection of 3 mm did not meet the expectations, compared to the reference specimens. this phenomenon seemed to be accidental. the recording of the test (fig. 6) also captured slight increment modified composite elastic modulus. similarly to micro-fibers, the behavior monitored using dic technique during four-point bending tests can be described as single crack and multi cracking in the case of samples reinforced with reference and plasma treated fibers, respectively. the results of dic analysis are shown in the fig. 4. 4.3. conclusion the main goal of this work was to enhance inter-phase interaction between cement matrix and polymer fiber reinforcement. the matrix was composed of portland cement (cem i 42.5 r, w/c ratio equal to 0.4). as the randomly dispersed and oriented reinforcement, two types of polypropylene fibers having 18µm and 80 µm in diameter were used. the interaction between the two materials was enhanced via fiber surfaces modification by low-pressure cool coupled oxygen plasma. a duration of fiber exposition to plasma was equal to 30 seconds. prismatic specimens having dimensions equal to 40×40×160 mm contained from the matrix and fibers (volume amount 2 %) were made and after 28 days of mixture curring and hardening subjected to four-point bending destructive tests. as an indirect indicator of both chemical and physical interaction between the two materials, the response (toughness and loadbearing capacity) of loaded specimens was monitored and evaluated. the results of samples reinforced with reference and treated fibers were mutually compared. the increase of load-bearing capacity of samples re14 vol. 13/2017 polymeric fiber reinforced cement composite inforced with modified fibers (both microand macrofibers) was shown. compare to reference mixtures, specimens containing modified fibers exhibited toughness increase in both linear and post-cracking phase. also, as for the plasma treated micro-fiber reinforced samples, significant increase of residual toughness (compared to reference specimens) was observed. similar findings were pointed out in the study examining the interaction between plasma modified pet (polyethylene terephthalate) and cement matrix [13]. dic analysis revealed a different behavior during four-point bending tests between samples containing reference and modified fibers. while the behavior can be described as single crack in the case of reference mixtures, the mixtures with modified fibers exhibited multi cracking behavior. this phenomenon clearly point to the required increase of the adhesion between treated fibers and the cement matrix. based on these findings, it can be pointed out, that polymer fiber modification via plasma treatment can be assumed as low-cost and very effective method to increase inter-phase interaction and mechanical properties of fiber reinforced cement composites. acknowledgements this work was financially supported by czech technical university in prague – sgs project sgs16/201/ohk1/3t/11 and by the czech science foundation research project 15-12420s. the authors would also like to thank to tereza horová for her help with this research. references [1] p.s. song, s. hwang, b.c. sheu, townsend t.g., strength properties of nylon and polypropylene fiber-reinforced concretes. cement and concrete research 35(8):1546–1550, 2005. doi: 10.1016/j.cemconres.2004.06.033 [2] n. banthia, r. gupta, influence of polypropylene fiber geometry on plastic shrinkage cracking in concrete, cement and concrete research 36(7): 1263–1267, 2006. doi: 10.1016/j.cemconres.2006.01.010 [3] s. kakooei, h. akil, m. jamshidi, j. rouhi, the effects of polypropylene fibers on the properties of reinforced concrete structures, construction and building material 27(1): 73–77, 2012. doi: 10.1016/j.conbuildmat.2011.08.015 [4] en 14889-2, fibers for concretepart 2: polymer fibers – definitions, specifications and conformity. [5] m. zeiml, d. leithner, r. lackner, h.a. mang, how do polypropylene fibers improve the spalling behavior of in-situ concrete?, cement and concrete research 36(5): 929–942, 2006. doi: 10.1016/j.cemconres.2005.12.018 [6] j. topič, j. bartoš, l. kopecký, j. fládr, z. prošek, j. trejbal, development of mechanical properties of a cement composite reinforced with synthetic fibers, trans tech publications ltd, 332–335, 2016. [7] a. peled, h. guttman, treatments of polypropylene fibres to optimize their reinforcing efficiency in cement composites, cement and concrete composites 14(4): 277–285, 1992. doi: 10.1016/0958-9465(92)90026-r [8] v. machovič, l. lapčák, l. borecká, m. lhotka, j. andertová, l. kopecký, l. mišková, microstructure of interfacial tranyition yone between pet fibres and cement paste, acta geodyn. geomater. 10(169): 121–127, 2013. doi: 10.13168/agg.2013.0012 [9] a.m. lópez-buendía, m.d. romero-sánchez, v. climent, c. guillem, surface treated polypropylene (pp) fibers for reinforced concrete, cement and concrete research 54: 29–35, 2013. doi: 10.1016/j.cemconres.2013.08.004 [10] j. trejbal, v. šmilauer, a. kromka, š. potocký, l. kopecký, wettability enhancement of polymeric and glass micro fiber reinforcement by plasma treatment, in: proceedings nanocon, brno, 2015. [11] j. trejbal, j. fládr, l. kopecký, utilization of plasma treated polymeric macro-fibers as reinforcement in concrete constructions, advanced materials research 1144: 70–75, 2016. doi: 10.4028/www.scientific.net/amr.1144.70 [12] b.h. yang et al., strength characteristics of modified polypropylene fiber and cement-reinforced loess, journal of central south university 24(3): 560–568, 2017. [13] j. trejbal et al., impact of surface plasma treatment on the performance of pet fiber reinforcement in cementitious composites, cement and concrete research 89: 276–287, 2016. doi: 10.1016/j.cemconres.2016.08.01 [14] b. pan et al., full-field strain measurement using a twodimensional savitzky-golay digital differentiator in digital image correlation, optical engineering 46: 1–10, 2007. doi: 10.1016/j.cemconres.2016.08.01 15 acta polytechnica ctu proceedings 13:11–15, 2017 1 introduction 2 materials 2.1 polymeric fibers 2.2 cement paste 3 experimental methods 3.1 plasma treatment 3.2 four-point bending test 3.3 digital image correlation 4 results 4.1 samples reinforced with micro-fibers 4.2 samples reinforced with macro-fibers 4.3 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0131 acta polytechnica ctu proceedings 15:131–136, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app strength of composite yarn under biaxial loading soňa valentováa, ∗, vladimír hrbeka, b, jan vorela, michal šejnohaa a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b institute of theoretical and applied mechanics as cr, v.v.i., prosecká 76, 190 00 prague 9, czech republic ∗ corresponding author: sona.valentova@fsv.cvut.cz abstract. the present paper is concerned with two types of periodic unit cell of composite yarn with different geometry but the same material properties. their macroscopic response under tensile and compressive loading in the transverse direction and their combination are plotted in the graphs. based on stress-strain curves the failure envelopes are constructed. a simple maximum stress criterion and linear softening law is used in the adopted progressive damage analysis to outline the softening part of stress-strain diagrams. finally the impact of selected representative volume element is observed through the comparison of results gained for both designed periodic unit cells in the microlevel, meaning the level of yarn. keywords: composite yarn, periodic unit cell, strength, damage analysis, failure envelope. 1. introduction demands on material properties increase as many different industry branches develop. therefore new materials, like textile composites become popular, especially for their variability, that consist for example in type of material used for fibers and matrix, type of weave, volume fraction of individual phases, curing temperature and so on. every mentioned factor influences the final macroscopic properties, so one of their big advantage is the possibility to adapt their final macroscopic properties. for their high specific strength, stiffness and temperature resistance they find use for example in automobile or aerospace industry. one of the most typical reinforcement are glass fibers, which belong to the most price favorable fibers and also have quite high modulus and strength. next commonly used fibers are carbon and basalt fibers. these are better in terms of mechanical properties. nowadays carbon fibers are still more often substituted by the basalt fibers thanks to their similar mechanical properties and lower price. for the possibility an extensive and mainly safe use of such material the knowledge of their behavior under different loading cases for both, in the elastic part and also after the damage occur, is necessary. as the experiments are complicated not only in terms of feasibility, but also from the time and financial point of view, some effective tool to obtain their macroscopic response numerically is desirable. the structure of braided composite is very complex, thus the appropriate simplification provided through the determination of representative volume element and subsequently construction of periodic unit cell, substituting its structure is important for the numerical analysis. in this paper, we adopt two types of periodic unit cell in the micro-scale, meaning the level of yarn, with attention concerned to the evaluation of macroscopic strength for unidirectional fiber composite under tensile and compressive on-axis loading in the transverse direction, see fig. 3 and fig. 4, and using their combination the failure envelope expressing the effective strengths in arbitrary transverse direction could be draw up. in case of textile composite, see the cross section and 3d periodic unit cell of such a composite in fig. 1, for obtaining its macroscopic properties, results from analysis introduced in this work will be applied to the meso-scale level where the effective properties of yarns are needful. moreover, the influence of elected periodic unit cell of yarn is investigated via the evaluation of effective elastic properties with the hand of finite element method subsequently compared with elastic properties gained by the mori-tanaka averaging scheme. 2. periodic unit cell of yarn design of periodic unit cell (puc) of yarn is based on the image analysis carried out on the cross sections of yarns, see fig. 2a, consisting of determination of geometrical parameters and then evaluation of volume fractions of individual phases. based on the knowledge of the inner geometry a representative volume element (rve) is determined, for more details see [1]. typical and the most simplified puc of yarn dealing with the idealized hexagonal distribution of fibers was designed, can be seen in fig. 3. moreover, 4-fiber puc of yarn with more complex rve was created, see fig. 4, to be able to verify the adequacy of designed puc. 131 http://dx.doi.org/10.14311/app.2018.15.0131 http://ojs.cvut.cz/ojs/index.php/app s. valentová, v. hrbek, j. vorel, m. šejnoha acta polytechnica ctu proceedings (a). (b). figure 1. (a) cross section of braided composite, (b) 3d-sepuc of composite. (a). (b). figure 2. (a) cross section of yarns with coordinate system, (b) example of mesh of yarns in composite. 3. failure criteria and damage model when loading of periodic unit cell representing yarn of braided composite we can distinguish two types of failure modes: • fiber failure mode • matrix failure mode in contrast with the direct stiffness reduction method, which simply assign zero stiffness to the integration point of mesh element, where the strength is reached, in adopted progressive damage model proposed by fang et al. [2, 3] lower individual elements of the stiffness matrix according to the magnitude of loading to reflect a material degradation with the evolution of cracks. then we can observe the macroscopic response also after the strength is exceeded. damage evolution is expressed as dj = χjfeq (χjeq −χjieq) χjeq(χ jf eq −χjieq) , j = 1t, 2t, 3t, 1c, 2c, 3c (1) in eq. 1 χjieq and χjfeq represents the equivalent displacement at the onset of failure and the equivalent displacement at the full damage, respectively.χjeq expresses equivalent displacement in the current loading step. the initial value is defined from the first load increment as χjieq = χjeq√ fj (2) where fj is jth failure criterion. as a simple maximum stress criterion is adopted in present work, following condition must be fulfilled to start developing the damage after the crack initializes. σii sij >= 1, i = 1, 2, 3; j = t,c (3) figure 3. 2-fiber puc puc-1. figure 4. 4-fiber puc puc-2. to define the displacement at the full damage in the nonlinear constitutive model, the fracture energy as the principal material property is indispensable. it can be written as∫ wfc 0 σdwc = ls ∫ fd 0 (�− �el)dσ = ls ∫ �f 0 σd� (4) the fracture energy is graphically expressed in fig. 5. (a). (b). figure 5. (a) stress-crack opening displacement diagram with linear softening law, (b) stress-strain diagram with softening. 4. numerical experiment numerical experiment was carried out on a basalt fiber composite with pyrolized polysiloxane matrix. as the material is heterogeneous, several indentation measurements were performed on both phases of the available sample, fibers and matrix. results from these measurements are listed in tab. 1. average results of elastic material properties partially taken from the experiment, partially from literature, see [4] and partially set according to the analogy with the carbon fibers, are summarized in tab. 1. young’s modulus was derived from indentation results based on contact mechanics principles of two colliding bodies, as described in [5, 6]. these results have already been presented in [7]. principles of the nanoindantation method are described in [5]. effective elastic properties of investigated composite summarized in tab. 3 were also investigated in order to 132 vol. 15/2018 strength of composite measurement red. mod. young’s mod. hardness contact depth number er [gpa] e [gpa] h [gpa] hc [nm] basalt fiber – longitudinal direction 1 67.81 ± 8.09 67.72 ± 7.68 5.27 ± 0.56 191.92 ± 6.55 2 75.95 ± 6.55 75.66 ± 6.21 5.95 ± 0.88 106.09 ± 8.91 3 76.85 ± 7.75 76.63 ± 7.35 5.71 ± 0.76 108.33 ± 7.98 4 66.17 ± 10.44 66.08 ± 9.93 5.49 ± 0.75 161.55 ± 12.22 5 71.76 ± 7.30 71.15 ± 6.92 5.79 ± 0.38 155.99 ± 5.62 6 61.12 ± 6.68 60.85 ± 6.33 5.28 ± 0.74 239.19 ± 18.71 average 69.94 ± 7.80 69.68 ± 7.40 5.58 ± 0.68 160.51 ± 10.00 basalt fiber – transverse direction 1 74.09 ± 6.22 73.65 ± 6.22 5.88 ± 1.17 107.60 ± 11.80 2 75.98 ± 3.79 75.69 ± 3.58 5.94 ± 0.54 105.58 ± 5.57 3 57.31 ± 3.58 56.85 ± 3.38 5.00 ± 0.55 116.33 ± 7.25 4 55.34 ± 10.55 54.80 ± 10.03 4.45 ± 1.16 126.64 ± 17.08 5 68.77 ± 8.09 68.65 ± 7.68 5.15 ± 0.54 114.39 ± 6.80 6 59.61 ± 11.70 59.26 ± 9.44 5.29 ± 1.22 117.95 ± 15.13 average 65.18 ± 7.32 64.82 ± 7.03 5.29 ± 1.19 114.52 ± 14.15 pyrolyzed polysiloxane matrix 1 67.57 ± 3.82 67.47 ± 3.61 7.29 ± 0.96 56.64 ± 4.25 2 56.45 ± 3.77 55.95 ± 3.56 5.44 ± 0.44 110.65 ± 4.90 3 60.20 ± 5.09 59.88 ± 4.82 6.01 ± 0.54 104.88 ± 5.21 4 56.70 ± 8.53 56.22 ± 8.10 5.63 ± 1.02 109.81 ± 10.39 5 58.39 ± 7.29 57.98 ± 6.91 5.68 ± 0.99 109.45 ± 11.72 6 55.82 ± 5.00 55.30 ± 4.73 6.23 ± 1.08 103.85 ± 10.21 average 59.19 ± 5.58 58.83 ± 5.29 6.05 ± 0.84 99.21 ± 7.78 table 1. indentation results of individual material phases. material e g ν gf st sc[gpa] [gpa] [–] [n/mm] [gpa] [gpa] fiber – longitudinal direction 69.68 28.10 0.40 2 –1.5 fiber – transverse direction 64.82 26.14 0.24 0.5 –1.5 pyrolyzed polysiloxane matrix 58.83 23.72 0.24 0.001 0.08 –0.16 table 2. considered parameters of individual materials. verify the adequacy of selected representative volume elements. they were obtained by the micromechanical mori-tanaka averaging scheme, which builds upon the knowledge of the shape and orientation of the reinforcement, material properties and volume fraction of individual phases. next, they were gained with the help of more complex first order homogenization approach, but on the other hand more numerical and time demanding method with the need of construction of puc. the elastic properties provided by the finite element method for two designed pucs illustrated in fig. 3 and fig. 4 agrees well with the mori-tanaka predictions. the main part of presented paper is numerical evaluation of strength of mentioned composite. such an analysis was performed in the transverse direction of both designed pucs. at first both the on-axis tensile and compressive loading and then their combinations necessary for construction of the failure envelope were applied. the failure envelope plotted in the fig. 10 illustrates the strength in the transverse direction of yarn. the shape perfectly corresponds to our presumption leading from the adopted simple maximum stress criterion mentioned above in eq. 3. the maximum and minimum values of reached effective stresses are marked and provided in the graphs for the on-axis loading directions. further and the last output of this study is evaluation of the nonlinear macroscopic response of the composite. for this purpose, stress-strain curves for both pucs were generated (fig. 6 to fig. 9). one for each on-axis transverse direction for tensile and compressive loading increments. 133 s. valentová, v. hrbek, j. vorel, m. šejnoha acta polytechnica ctu proceedings model type e11 e22 e33 g23 g31 g12 cf[gpa] [gpa] [gpa] [gpa] [gpa] [gpa] [–] sepuc-1 fem 62.96 62.97 65.86 26.46 26.48 23.5 0.62 sepuc-2 fem 62.98 62.96 65.86 26.48 26.46 23.5 0.62 mori-tanaka method 62.97 62.97 65.88 26.47 26.47 23.5 0.62 table 3. effective elastic moduli. figure 6. stress-strain curves for on-axis tensile loading in the x-direction. figure 7. stress-strain curves for on-axis tensile loading in the y-direction. evidently the macroscopic response gained for both pucs gives almost the same results not only in the elastic part but also for the nonlinear behavior with just negligible differences. 5. conclusion the presented paper was concerned with two types of periodic unit cell of composite yarn with different representative volume element but the same material properties. at first their effective elastic properties were obtained both by the mori-tanaka averaging scheme and finite element method. based upon gained figure 8. stress-strain curves for on-axis compressive loading in the x-direction. figure 9. stress-strain curves for on-axis compressive loading in the y-direction. results the adequacy of the designed puc could be verified. in the next step the macroscopic elastic and nonlinear response for tensile and compressive loading in the transverse on-axis direction were investigated and plotted in the diagrams. on the basis of the presented methodology we can observe macroscopic behavior of composite after cracks initialize and damage starts to develop. a simple maximum stress criterion and linear softening law was adopted in this study to outline the nonlinear part of stress-strain diagrams. based on the stress-strain curves derived from the biaxial loading, the failure envelopes were constructed. re134 vol. 15/2018 strength of composite figure 10. failure envelope for biaxial loading in the transverse direction. alized computational analysis shows the macroscopic strength of investigated composite for the input parameters summarized in tab. 3. differences in failure envelopes (fig. 10) and in stress-strain curves raised for both examined pucs (fig. 6 to fig. 9) are almost imperceptible and could be caused by the unequal generated mesh density. consequently we can afford to consider the disparities such insignificant that the simplification of yarn through the designed puc is perfectly acceptable. the model was tested on mentioned composite but with incomplete knowledge of needed input material parameters. in the future work we will pay our attention to another numerical and nanoindentation experiments to get necessary parameters to accomplish more accurate analysis on micro-scale and continue with numerical predictions of composite with more complex structure, e.g. illustrated in fig. 1. list of symbols ν poisson ratio [–] er reduced modulus [gpa] e young’s modulus [gpa] h hardness [gpa] hc contact depth [nm] g shear modulus [gpa] gf fracture energy [n/mm] st critical strength in tension [gpa] sc critical strength in compression [gpa] eij homogenized elastic modulus [gpa] gij homogenized shear modulus [gpa] dj damage parameter [–] χeq equivalent displacement [m] σij stress [gpa] �ij strain [–] acknowledgements the financial support provided by the sgs project with the application registered under the no. sgs18/037/ohk1/1t/11 is gratefully acknowledged. references [1] j. zeman, m. šejnoha. from random microstructures to representative volume elements. modelling and simulation in materials science and engineering 15(4):325–335, 2007. doi:10.1088/0965-0393/15/4/s01. [2] g. d. fang, j. liang, q. lu. investigation on the compressive properties of the three dimensional four-directional braided composites. composite structures 93(2):392–405, 20011. doi:10.1016/j.compstruct.2010.09.002. [3] g. d. fang, j. liang, b. l. wang. progressive damage and nonlinear analysis of 3d four-directional braided composites under unidirectional tension. composite structures 89(1):126–133, 20009. doi:10.1016/j.compstruct.2008.07.016. 135 http://dx.doi.org/10.1088/0965-0393/15/4/s01 http://dx.doi.org/10.1016/j.compstruct.2010.09.002 http://dx.doi.org/10.1016/j.compstruct.2008.07.016 s. valentová, v. hrbek, j. vorel, m. šejnoha acta polytechnica ctu proceedings [4] m. černý, p. glogar, z. sucharda. mechanical properties of basalt fiber reinforced composites prepared by partial pyrolysis of a polymer precursor. journal of composite materials 43(9):1109–1120, 2009. doi:10.1177/0021998308097732. [5] w. c. oliver, g. m. pharr. measurement of hardness and elastic modulus by instrumented indentation: advances in understanding and refinements to methodology. journal of material research 19(1):3–20, 2004. doi:10.1557/jmr.2004.0002. [6] h. hertz. on the contact of rigid elastic solids. hertz’s miscellaneous papers p. chap. 6, 1896. original: über die berührung fester elastischer körper, journal für die reine und angewandte mathematik, 92: 156–171, 1881. [7] s. valentová, v. hrbek, m. šejnoha. evaluation of effective properties of basalt textile reinforced ceramic matrix composite. acta polytechnica ctu proceedings 13:142–147, 2017. doi:10.14311/app.2017.13.0142. 136 http://dx.doi.org/10.1177/0021998308097732 http://dx.doi.org/10.1557/jmr.2004.0002 http://dx.doi.org/10.14311/app.2017.13.0142 acta polytechnica ctu proceedings 15:131–136, 2018 1 introduction 2 periodic unit cell of yarn 3 failure criteria and damage model 4 numerical experiment 5 conclusion list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0094 acta polytechnica ctu proceedings 15:94–98, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app microscopic and phase analysis of cement paste containing waste micronized marble powder zdeněk prošeka, b, ∗, pavel tesáreka, jan trejbala a czech technical university in prague, faculty of civil engineering, thakurova 7, 166 29 prague 6, czech republic b czech technical university in prague, uceeb, trinecka 1024, 273 43 bustehrad, czech republic ∗ corresponding author: zdenek.prosek@fsv.cvut.cz abstract. the article focuses on the impact of waste marble powder on resulting cement composites. we investigate of influence waste marble powder on the hydration process and hence the resulting phase composition. the hydration process was investigated using calorimetry and phase composition of the resulting composite by electron microscopy. the results are compared with a reference sample composed of portland cement. keywords: micronized marble powder, marble sludge, cement composites, calorimetry, elemental analysis. 1. introduction the great demand for marble in export markets has led to increased his production. an important factor is the amount of mined stone, regardless of the quantity of waste. production lines for mining and stone processing produce two types of waste. the first one is in the form of crushed stone that arises by removing marble blocks from the line due to their bad shape or when the blocks are broken or cracked. the production of this waste is great, but its application is so effective in the form of aggregates and gravel to the concrete that it does not create any ecological risks. the second form of waste is marble sludge and dust, which occurs during the mining, crushing and cutting of marble blocks or during the final grinding and polishing process. the amount of waste produced is estimated, for example, for the egypt region at one million tonnes per year [1]. the amount of waste produced is not only associated with less advanced countries such as the middle east but also with europe, where, for example, in portugal, the amount of waste generated is estimated at 600 thousand tonnes per year [2]. an approximately 10 to 20% of the waste is further processed and used [3]. the most common use is in road infrastructure, as asphalt filler or underlayer [4]. the remaining unused waste is stored in the vicinity of the factories, regardless of their possible use, thus causing serious environmental problems. for these reasons, it is important to create an application of these wastes, which would result in improved safety and health protection by lowering dust particles in the air, lowering the cost of storage, transport, and waste handling, and generating potential revenue [5]. one option is to use waste as a partial substitute for portland cement, where recycling material could have a positive effect on resulting properties of cement composites. the resulting properties of the cement matrix are directly dependent on the composition. the resulting cement matrix is composed of a plurality of phases. the largest proportion of the phases has the c-s-h gel (hydrosilicate), which is furthermore formed by tobermorite, afwillite, xonotlite and other still unknown phases. furthermore, hydroaluminate, gypsum, ettringite, ferrite hydrate and portlandite (ch crystals) appear in the cement matrix [6]. they are several works that solve the possibility of replacing cement with marble sludge, and the result is two contradictory findings. the first statement is described in a study by demirel [7] at the firat technical university in turkey. he found using an electron beam microscope that when the marble sludge was added to the cement composite, the ch crystal size was changed. mixtures without marble sludge had a form of large ch crystals. while in a mixture with marble sludge had very small ch crystals and it was produced across all the matrix. this effect is explained by the fact that fine sludge grains act in the cement composite as nucleation centers for the growth of ch crystals and accelerate the hydration of silicate and aluminate phases [8]. the second claim is described in the work of aliabdo et al. [9]. they examined the morphology of the internal structures of the cement slurry with and without marble sludge by using a scanning electron microscope and they found that, is not distinctly different from each other, meaning that the marble sludge has no noticeable role during the hydration process and therefore only acts as an inert filler in the new product. with the development of the technique, new possibilities arise to confirm or rebut possible claims about the inclusion of marble sludge during hydration. one possibility is the treatment of marble sludge using high-speed milling [10], where the resulting micronized marble powder has a higher density and therefore a higher reactivity [11]. this recycling method has already been successfully used 94 http://dx.doi.org/10.14311/app.2018.15.0094 http://ojs.cvut.cz/ojs/index.php/app vol. 15/2018 microscopic and phase analysis ... figure 1. the image of micronized marble powder, magnified 100×, bse detector. for concrete recycling [12]. 2. materials and samples portland cement cem i 42.5 r was chosen as reference material from českomoravský beton, a.s. with a factory in radotín. it is a cement used for high strength concrete and mortars and plasters. the main ingredient is portland clinker with a proportion of 95 to 100 wt. %. the remaining amount of weight is the complementary components. the marble sludge that was created during the cutting and polishing of limestone marble was chosen for investigating the possibility of replacing cement with micronized mechanically activated marble powder. the marble sludge comes from an industrial area located on the west bank. the marble sludge was lumpy and contained big grains. the marble sludge was micronized on high-speed mills in collaboration with lavaris for improved workability, improved marble sludge properties, and increased surface area. the high-speed mill type sbd 600 with a power of 2×30kw with patented cutting edges was used for this purpose. two grinding modes and two types of grinding elements were used. in the first case, it was used element in the form of patented 400 mm diameter pins, and in the second case patented teeth with a diameter of 150 mm. in both cases, the grinding speed was set to 300 m/s. the result of high-speed grinding was micronized marble powder. two sets of samples were created for research. the first set was used for calorimetry testing, consisting of cement pastes with varying amounts of micronized marble powder. mixture m0 served as a reference and contained only cement. mixtures m5 to m50 contained portland cement and micronized marble powder in 5, 10, 15 and 50 wt. % (table 1). the water ratio, i.e., the weight of water to the total dry mass mixture, was 0.45. the second set was used for electron microscopy testing. two microscopic samples were created for the examination of the materials at the nano/micro level. the first sample was made by mixing micronized marble with epoxy resin and served to examine the size and shape of the grains and to analyze the recyclate itself. the second sample was made of a hardened cement paste which was made up of 50 wt. % portland cement and 50 wt. % recycled micronized marble powder. it was sample m50 from calorimetric measurements. before creating the cuts, it was necessary to fill the pore space with a suitable substance so as not to release the grains of materials during the preparation of the samples. the vacuum impregnation method for epoxy resin samples (epofix kit) was used for this purpose. the polished sections were prepared on struers’ tegramin machine. the samples were ground and polished in several steps to achieve the best surface quality of the samples. a different preparation procedure was performed for each sample due to different material. the first sample, the recycled micronized marble powder is homogeneous and does not react with water, so it is possible to use silica paper and water-based lubricant. in the first step, silica paper with a grain size of 220 grains/cm2 was used to remove the greatest unevenness after cutting. in the following steps, finer silica paper was used: 500 grains/cm2, 1200 grains/cm2, 2000 grains/cm2 and 4000 grains/cm2. each step lasted 2 minutes, and the pressure on the sample was 5 n. in the last step the samples were polished by hand using a 1/4 microns emulsion containing the nanodiamonds for 1 minute. the second sample is made up of several phases, so we used diamond plates that do not affect this inhomogeneity and does not occur relief in the grinding and polishing process. also, the cement reacts with water, so an alcohol-based lubricant has been used. md-piano 500 was used in the first step (corresponding 500 grain / cm2 silica paper roughness), and the next steps were md-piano 1200, md-piano 2000 and md-piano 4000. the grinding and polishing time was several hours because polishing and grinding were performed without pressure. finally, both samples were sputtered with a carbon layer at 30 nm to improve the required surface conductivity of the sample and improve the quality of the analysis. 3. experimental methods and results the development of hydration heat was measured to measure the effect of micronized marble powder on the hydration process of portland cement. measurement of the development of the hydration heat was performed on a tam air calorimeter. tam air is an isothermal calorimeter for accurately determining the heat flow (production and consumption) and the amount of heat produced. a total of 10 chambers were used for the measurements (two chambers for each mixture). the mixtures were tested for 7-day measurement of the development of heat flow at a constant temperature of 20 ◦c. the mixtures were stored in plastic closable containers which were described, weighed and eventually compacted before insertion into the calorimeter. each container contained approximately 25-38 g of the mixture. the output from the calorimetric measurement was the heat flow in w. 95 z. prošek, p. tesárek, j. trejbal acta polytechnica ctu proceedings mixtures replacing cement cement (cem i 42.5 r) [g] micronized marble powder [g] water [g] m0 0 wt. % 80 0 36 m5 5 wt. % 76 4 36 m10 10 wt. % 72 8 36 m15 15 wt. % 68 12 36 m50 50 wt. % 40 40 36 table 1. composition of the individual set/mixtures. element weight wt. [%] wt. standard deviation [%] atomic [%] c 16.93 0.33 30.06 o 32.29 0.50 43.04 mg 0.32 0.10 0.04 ca 50.47 0.41 26.86 total 100.00 100.00 table 2. evaluation of elemental analysis. figure 2. curve of heat evolution for tested materials. the results from the calorimetric measurement were converted to 1 g of cement based on the masses of the mixtures in the container. the feg sem merlin zeiss scanning electron microscope was used for microscopic structural and elemental analysis, which is located in the laboratory of electron microscopy and microanalysis at the university center of energy efficient buildings. qualitative and quantitative analysis of the chemical composition of the samples was performed using an x-ray microanalysis, namely an energy spectrometer (eds) by oxford instruments. the point id method was used in the first sample, where only selected areas (areas with marble limestone) were tested. the second sample was examined using line scans. the method best describes the possible effect of micronized marble powder on the final product. the microscope setting was such that the accelerating voltage was 10 kv, the current 1 na, the working table distance 8.5 mm, the measuring time of one point 50 µs and the resolution 1024 px. the composition of the recycled micronized marble powder can be determined from the results from the microscopic elemental analysis. figure 1. shows the areas under investigation, and the results from the microscopic elemental analysis are presented in table 2. the results are slightly distorted due to the need for sputtering the samples with a 30 nm carbon layer. it can be stated that about 99% of the waste material is made up of caco3 (calcite) and the remaining 1% is of minor constituents, the most important of which is mgco3. from the pictures taken during the microscopic analysis, it is clear that the grains consist 96 vol. 15/2018 microscopic and phase analysis ... figure 3. the image of cement paste containing micronized marble powder, magnified 500×. figure 4. concentration of calcium and silicon in line scan. of one phase, they have a sharp and oblong shape. the calorimetric measurement (figure 2.) shows the effect of micronized marble powder on the developing of hydration heat of the mixture. thanks to the conversion of the heat flow to 1 g of cement, the differences in the hydration process were highlighted. for mixtures with micronized marble powder, it can be observed that after 20 hours of mixing, there was a slight increase in the heat flow, probably due to the hydration of c3s. the difference in heat flow is apparent up to approximately three days (m50). further, the course of heat flow is similar, i.e., micronized marble powder has no significant effect on c2s hydration. this effect can be caused either by marble powder or the addition of finely ground particles. for this reason, we have proceeded to microscopic elemental analysis. figure 3 shows the examined area for which four line scans were created. a backscatter electron detector was used to display the phase contrast information. the result was similar for each scan, so one was chosen as representative, on which the results were described (figure 3). figure 4 shows the weight distribution of calcium and silicon and also their ratio (ca / si). emphasis was placed on these elements because cem i 42.5 r radotin has a high content of c3s (tricalcium silicate) and c2s (dicalcium silicate) and low content of c3a (tricalcium aluminate) and c4af (tetracalciumaluminoferite), i.e., the resulting product is mainly composed of these elements. at the point of occurrence of the c-s-h gel, that is, the phases resulting from the hydration and hydrolysis of c3s and c2s, the amount of calcium and silicon affect the resulting phases, where the greatest proportion of calcium has xonotlite and the smallest afwillit. overall, the identification and precise composition and structure of hydro silicates (hydro silicate gels) is difficult and therefore only the ca/si ratio is indicated. the resulting c-s-h gel can be chemically described 97 z. prošek, p. tesárek, j. trejbal acta polytechnica ctu proceedings as follows: xcao · ysio2 · zh2o, (1) where the high density c-s-h gel is in range: x = 0.5 − 1.5, y = 1, z = 0.5 − 2.5, and low density c-s-h gel x = 1.5 − 2, y = 1, z = 1 − 4. if we look at the weight of calcium (figure 4), we see that there is a greater concentration of calcium at the site of the grain. a large amount of calcium in the grain is due to its composition because it is mostly composed of calcium, i.e., caco3. ca/si ratio is increased up to a distance of about 50 microns from the grain, mainly due to the reduction in the amount of silicon, followed by a further 30 microns of gradual normalization of the values. from the results, it can be stated that another phase is preferred in the vicinity of the grains. this phase could be a low-density c-s-h gel. 4. conclusions this work focuses on the effect of micronized marble powder as a partial substitute for portland cement. the resulting cement pastes contained portland cement and micronized marble powder in 5, 10, 15 and 50 wt. %. based on the results, it can be concluded that: • the high-speed milling process generates grains that are angular and oblong. • calorimetry results showed that after 20 hours of mixing, there was a slight increase in heat flow caused by c3s hydration, the heat flow was the same, so the marble powder did not have a significant effect on the hydration of c2s. • from the results of the elemental analysis, it can be stated that the creation of one phase is preferred in the vicinity of the grains. this effect is probably because around the grain of marble is less free water needed to hydration of the cement. this phase could be a low-density c-s-h gel. the results of the work are in contradiction with the statement of demirel [7] and aliabdo [9], because marble powder during the hydration process is not completely inert, and smaller ch crystals do not appear in the surroundings of grains. in the future, the research will focus on confirming the theory of creating a low-density c-s-h gel around the grain of marble. for this reason, further work will be devoted to micromechanics around marble grains using nanoindentation. acknowledgements this outcome was supported by the czech technical university in prague under no. sgs16/201/ohk1/3t/11, by the czech science foundation research projects gacr 17-06771s and by the ministry of education, youth and sports within national sustainability programme i, project no. lo1605. the authors also thank the center for nanotechnology in civil engineering at the faculty of civil engineering, czech technical university in prague. references [1] a. mashaly, b. shalaby, m. el-hefnawi. characterization of the marble sludge of the shaq el thoaban industrial zone, egypt and its compatibility for various recycling applications. australian journal of basic and applied sciences 6(3):153–161, 2012. [2] l. lopes, r. martins. marbles from portugal. naturstein, online publication, naturalstone-online com 2012. [3] m. karaşahin, s. terzi. evaluation of marble waste dust in the mixture of asphaltic concrete. construction and building materials 21(3):616–620, 2007. [4] p. mehta, v. k. mehta. waste generation and minimization: a study of marble mines of rajsmand. international journal of informative & futuristic research 2(09):3049–3058, 2015. [5] m. el-gammal, m. ibrahim, e. badr, et al. health risk assessment of marble dust at marble workshops. nature and science 9(11):144–154, 2011. [6] t. c. powers. structure and physical properties of hardened portland cement paste. journal of the american ceramic society 41(1):1–6, 1958. [7] b. demirel. the effect of the using waste marble dust as fine sand on the mechanical properties of the concrete. international journal of physical sciences 5(9):1372–1380, 2010. [8] r. hela. příměsi do betonu / concrete admixtures . [9] a. a. aliabdo, a. e. m. a. elmoaty, e. m. auda. re-use of waste marble dust in the production of cement and concrete. construction and building materials 50:28–41, 2014. [10] z. prošek, a. osvald, j. topič, et al. influence of recycled concrete composition on its elastic stiffness. in key engineering materials, vol. 677, pp. 288–291. trans tech publ, 2016. [11] z. prošek, j. trejbal, j. topič, et al. utilization of the waste from the marble industry for application in transport infrastructure: mechanical properties of cement pastes. in iop conference series: materials science and engineering, vol. 236, p. 012092. iop publishing, 2017. [12] j. topič, z. prošek. properties and microstructure of cement paste including recycled concrete powder. acta polytechnica 57(1):49–57, 2017. 98 acta polytechnica ctu proceedings 15:94–98, 2018 1 introduction 2 materials and samples 3 experimental methods and results 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0001 acta polytechnica ctu proceedings 15:1–5, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app bond connections of high performance concrete (hpc) and steel: "experimental verification of selection glued for two different surfaces of steel" tomáš blažeka, b, ∗, jakub řepkaa, b, tomáš vlacha, b a czech technical university in prague, faculty of civil engineering, thakurova 7, 166 29 prague 6, czech republic b czech technical university in prague, uceeb, trinecka 1024, 273 43 bustehrad, czech republic ∗ corresponding author: tomas.blazek@fsv.cvut.cz abstract. this article deals with the test of selected adhesives for glued connection of high performance concrete and steel plates with various surface treatments. results of this experiment can be used as a basis for the possible anchoring of concrete facade elements to the steel grate, the possible bonding of steel anchors on the surface of concrete facade elements. high performance concrete is a concrete that exceeds the compressive strength 100 mpa after 28 days. the second adherent is the uncoated steel and the steel with zinc coating as a second variant. samples were tested in two different ways. the first experiment verified the single tensile strength of bond connection. the second test verified the shear strength, which is very important for connection designing. keywords: hpc, concrete, steel, joint, connection, glue, epoxy resin. 1. introduction high performance concrete (hpc) has been increasingly used already also in the czech republic both in structural such as bridges and other structures such as indoor and outdoor furniture or facade elements. research at the department of building structures, faulty of civil engineering ctu in prague is much focused on the development of lightweight concrete facade elements [1, 2]. due to the corrosion of traditional steel reinforcement is more and more used alternative reinforcement in facade elements such as non-steel fibres, impregnated and non-impregnated technical textiles [3, 4]. it leads to the next lightening of elements hand in hand with development of hpc [5, 6]. their anchoring to the building structure is an essential part of the research. due to their lower weight less anchorage capacity is required and it is getting interested also the bonding in addition to the traditional mechanical anchors [7, 8]. the existing facade elements are usually anchored using mechanical anchors of various shapes with different mechanisms and systems for rectification. these mechanisms can be visible or hidden and sometimes affect the view on the building envelope. testing the mechanical anchors of the circumferential envelopes is addressed in more detail for example in etag 034 [9]. the aim of this work was selection of suitable adhesives commonly available on the market and their experimental verification on the connection of two basic materials. it was a high performance concrete (hpc) element representing for example facade element and steel plate representing steel grate. both materials hpc and steel have almost zero porosity and minimal water absorption. these are more demanding conditions for bonding. adhesives are usually tested for one-stage stress test blunt joint according to the csn en 15 870 [10] standard and they are also usually tested for shear stress according to the csn en 1994-1-1 [11] standard. the shear test is not exactly according to the standard. dimensions for presented experiment were approximately adapted to the dimensions of the facade elements and to the possibilities of the laboratory testing equipment. finally four adhesives were tested on two selected commonly used types of steel adhesion finishes, black steel without any surface treatment and heat-galvanized adherent [12]. 2. materials used for experiment hpc developed at the faulty of civil engineering ctu in prague (fce ctu) from local sources of raw materials has been already used for many applications and experiments. the same hpc has been also used for presented experiments in this paper. mix design of used fine-grain self-compacting hpc is presented in table 1. water cement ratio was 0.25 and water binder ratio was only 0.20. compressive strength of used hpc was 113 mpa on cubes with edge length 100 mm according to csn en 12 390-3 standard. flexural strength (three point bending test) is 13.2 mpa on prisms 40 × 40 × 160 mm according to csn en 101511 standard. in the past on the same mixture also direct uniaxial tensile strength has been tested and it is 6.7 mpa on dog-bones with cross sectional area 30 × 30 mm according to csn 73 1318 standard and static modulus of elasticity in compression is 49.3 gpa 1 http://dx.doi.org/10.14311/app.2018.15.0001 http://ojs.cvut.cz/ojs/index.php/app t. blažek, j. řepka, t. vlach acta polytechnica ctu proceedings mix content kg/m3 cement i 42.5r (č.m. cement, mokrá) 680 technical silica sand 960 silica flour (ground quartz) 325 silica fume (microsilica) 175 superplasticizers 29 water 171 total 2,340 table 1. hpc mix design. on prisms 100 × 100 × 400 mm according to csn iso 6784 standard. two often used types of steel adhesion finishes were selected, black steel without any surface treatment and heat-galvanized adherent. the steel type was s235. in the experiment, four selected typical adhesives commonly available on the market were tested: mamut crystal glue, merbenit sf50, m3 and sika floor 156 – epoxy resin. the joint was experimentally subjected to tensile and shear stress and the mechanical behaviour of the adhesives was monitored. measured values were the crosshead displacement of testing machine, force and time of loading process. 3. experiments the basic hpc specimen dimensions for the tensile test were selected on the 100 × 25 × 25 mm. the hpc adherent part was glued to the steel part with a dimension of 50 × 25 × 25 mm. logically the fully glued cross sectional area was 25 × 25 mm. all selected types of glues were applied on this whole area to achieve the ideal source just after 28 days of hpc hardening from the concreting. described tensile test design and view on the real specimen in testing machine is presented in fig. 1. specimens after one week maturing were tested for the uniaxial tensile test with the calculated time until the failure 60 s ± 20 s. the displacement load speed was calculated on the 0.5 mm/min and it was used for all tensile specimens. the hpc adherent for the shear test were designed on 100 × 60 × 40 mm. dimensions of both steel adherents were similar 100 × 60 mm with different thickness 20 mm. the glued area in contact zone was 60 × 50 mm, but the specimens for shear test had two glued area because of required testing symmetry. it eliminated other undesirable stresses in the glued area. described shear test design and view on specimen in testing machine after the testing procedure is presented in fig. 2. the displacement load speed was calculated on the 0.2 mm/min and also in the shear test was same for all groups of specimens. 4. results of the tensile test the goal of this tensile test was the determination of maximal tensile stress until failure of the adhesive complex consisting of hpc adherence, steel adherence and adhesive. the stress values were calculated using critical maximum breaking force and the bonded area approximately 25 × 25 mm. the tensile stress dependence on the displacement of testing machine crosshead is presented in fig. 3 for all adhesives and also for both selected surface treatments of the metal adherence. presented curves are the average curves of three specimens for all selected types. the stress development was almost linear until the maximum value was reached. the description of stress development can be divided into three similar typical areas for all tested specimens. the first is the area before the linear trend of curves. this part was nonlinear due to the gradual activation and pre-stressing of all elements in testing machine. the second area is with linear trend of stress increasing on the cross-beam until the moment of maximal force was reached. different slope of curves suggests different stiffness of adhesives. it does not have to be the rule, but with higher adhesives stiffness also higher strengths have been achieved. epoxy resin sika floor 156 achieved the best results, the damage point was in the hpc part. commonly used construction adhesives in praxis on other chemical bases compared to the epoxy resin exhibited worse tensile properties. the third area is the part of curves behind the point of maximal force. for adhesives extruded from the tube (constructions or mounting adhesives), which has let’s say a rubbery soft character, this part shows considerable ductility. and it is a negative option of epoxy resin, high values of tensile stress, but fragile material behaviour. epoxy resin in general exhibits a sudden rupture, after the point of maximal force the epoxy resin is not able to transfer any load as visible in fig. 3. 5. results of the shear test in this part of the experiment the aim was to determine the maximum shear force and calculate the maximum shear stress before the failure of specimens. the shear stress was calculated based on the measured maximum force before the breaking point and the bonded two surfaces of the joint with cross sectional area approximately 60 × 50 mm. the dependence of the stress on the displacement is presented in fig. 4 in the form of graph, taking into account the surface treatment of the metal adherence similarly like in the case of previous described tensile test. presented curves are also the average curves of three specimens for all selected types. the description of stress development can be divided into three similar typical areas for all tested specimens similarly with previous described tensile test. also the mechanical behavior of the joint is very similar. epoxy resin reached the highest shear strength, but its behavior is fragile. in the moment when the maximum force in the contact area is reached, the 2 vol. 15/2018 bond connections of hpc and steel figure 1. design of tensile test: 1. steel sleeve, 2. steel adherent, 3. selective glue, 4. hpc adherent, 5. steel sleeve, 6. construction glue. view on the real specimen in testing machine. figure 2. results of the tensile test. joint is unable to transfer any load. in contrast, other softer adhesives have a distinct plastic part of deformation and their behavior is more ductile. however, the shear strength is lower compared to the epoxy resin sika floor 156. metal adherent without any surface treatment achieves better results compared to the metal adherent with zinc surface treatment. only in the case of merb glue opposite result were achieved. may be the surface of non-treated steel parts was incorrectly cleaned by technical gasoline or the surface was dusty or wet also in comparison with results of the tensile test. the amount of resin in connected part and the method of resin application could have a significant effect on the results. epoxy resin sika floor 156 and epoxy resins in general are two-component adhesives, where curing occurs when the both components are mixed together in the correct ratio. but in the cases of other adhesives (variants in tubes merb, 3m and mamut resin) reaction occurs with the access of air to the adhesive. therefore, manufacturers recommend that the glue in tubes should be applied in the caterpillars side by side. this procedure provide access of air along the caterpillars and good process of hardening. this procedure was followed, but it was very difficult in praxis estimate how the caterpillars should be apart form others. on several samples, the caterpillars were joined together along the length of the glued surfaces and adequate access of air was not ensured. for a shear test with a 50 × 60 mm of glued surfaces, the area was so large that the adhesive was not well hardened in 3 t. blažek, j. řepka, t. vlach acta polytechnica ctu proceedings figure 3. addiction stress – displacement, tensile stress. figure 4. addiction stress – displacement, shear stress. the middle of the sample in these cases. also use of hpc, which is waterproof, vapor-proof and airtight and does not allow access to air, is also of no effect. 6. conclusion the results of the tensile and shear tests correspond to each other. the best results with respect to strength values was achieved with epoxy two component resin sika floor 156. type of failure was for all samples was lost of adhesion on the side of hpc adherence. all other selected adhesives with soft rubber character (mamut, merbenit and m3) had a cohesive failure. it means that after the moment when the maximal force in the contact area was reached, the joint was able to transfer load with lower values of tensile stress. two selected commonly used types of steel adhesion finishes, black steel without any surface treatment and heat-galvanized, reached similar results. metal adherent without any surface treatment achieves may be little bit better average results compared to the metal adherent with zinc surface treatment. only in the case of merb glue opposite result were achieved because the surface of non-treated steel parts was incorrectly cleaned and prepared before the adhesive application. especially for soft rubber glues of lower strength (mamut, m3) the treatment of steel adherence is extremely important before adhesives are applied. the 4 vol. 15/2018 bond connections of hpc and steel manufacturer’s technical sheet must be always strictly observed. black steel is naturally characterized by a local occurrence of oxidized iron (scorch) which adversely affects the overall strength of the joint. the method of failure for presented adhesives was more or less cohesive with the visible small residues of oxidized iron. 7. summary the aim of this work was to verify the joining possibility of high performance concrete and steel adherence plates using adhesives. four commonly available selected adhesives were experimentally tested. the glued joint was verified in the tensile and shear strengths according to the relevant standards. the results show that the highest values were achieved using an epoxy resin based adhesive for both performed testing methods. the observation also showed a negative effect of the oxidized iron at the surface on the shear strength of the complex. tension strength is not fundamentally affected by this fact. this thesis does not solve the durability of used connection and its fire resistance, these propertieswill be verified in follow-up research. acknowledgements this outcome has been achieved with the financial support of the ta cr 11124 311 3111702a124 subtle concrete furniture and small structures for the railways stations and also has been supported by student grant competition, sgs18/108/ohk1/2t/11 – environmental aspects of high performance cement composites and concrete with recycled aggregate including their durability and service life. the authors would like to acknowledge with gratitude all the financial help that was provided in order to support this original research. references [1] t. bittner, p. bouška, m. kostelecká, et al. experimental investigation of mechanical properties of textile glass reinforcement. applied mechanics & materials 732, 2014. [2] a. chira, a. kumar, t. vlach, et al. textile-reinforced concrete facade panels with rigid foam core prisms. journal of sandwich structures & materials 2(18):200–214, 2016. [3] l. laiblova, t. vlach, a. chira, et al. technical textiles as an innovative material for reinforcing of elements from high performance concretes (hpc). advanced materials research 1054, 2014. [4] l. laiblová, t. vlach, ženíšek m. využití technických textilií jako alternativní výztuže do subtilních fasádních panelů z vysokohodnotného betonu. sympozium společnosti pro techniku prostředí pp. 155–160, 2015. [5] p. reiterman, m. jogl, v. baumelt, et al. development and mix design of hpc and uhpfrc. advanced materials research 982, 2014. [6] p. reiterman, o. holcapek, m. jogl, et al. long-term resistance of hpc to action of deicing salts. aip conference proceedings p. 090007, 2014. [7] etag 001. guideline for european technical approval of metal anchors for use in concrete, 2000. [8] p. bernardin. spojování kompozitních materiálů s kovem, 2016. disertační práce, zču v plzni. [9] etag 034. guideline for european technical approval of kits for external wall claddings, 2012. [10] čsn en 15 870. adhesives determination of tensile strength of butt joints, 2009. [11] čsn en 1994-1-1. eurocode 4: design of composite steel and concrete structures part 1-1: general rules and rules for buildings, 2006. [12] m. bajer, j. barnat. the glue–concrete interface of bonded anchors. construction and building materials 34:267–274, 2012. 5 acta polytechnica ctu proceedings 15:1–5, 2018 1 introduction 2 materials used for experiment 3 experiments 4 results of the tensile test 5 results of the shear test 6 conclusion 7 summary acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.21.0024 acta polytechnica ctu proceedings 21:24–27, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app glass waste powder utilization in high performance concrete diana mariakováa, b, ∗, tomáš vlacha, b, tereza pavlůa, b a czech technical university in prague, faculty of civil engineering, thákurova 2077/7, 166 29 prague 6, czech republic b czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: diana.mariakova@fsv.cvut.cz abstract. this paper deals with investigation of high-performance concrete (hpc) with full replacement of the silica powder by the waste glass powder. the silica powder was replaced by two types of the waste glass powder, originated from different sources (waste glass powder from grinding jewelry and milling of municipal waste glass). the properties of the waste glass powder were examined and compared with the silica powder. the mechanical and durability properties of three hpc mixtures were experimentally verified. the bulk density, flexural strength, and compressive strength were tested on beams 40 × 40 × 160 mm at age 28 and 60 days and after 0, 25, 50, 75 and 100 freeze-thaw cycles. there were observed slightly worse properties of mixtures with the waste glass powder in comparison with reference hpc. keywords: waste glass powder, high performance concrete, durability, mechanical properties. 1. introduction high-performance concrete (hpc) becomes very popular. hpc has been mostly used in the precast industry, for pedestrian footbridges. nowadays, hpc is used in buildings structures (skyscrapers) or in transportation structures (bridges) [1]. due to its excellent mechanical properties the hpc structures could be more subtle. this could lead to the raw materials saving. the properties of hpc are suitable in many ways such as mechanical and durability properties. hpc is also possible to use in aggressive environment with aggressive substances such as sulphates or chlorides. permeability of hpc is also very low [2]. there is also the risk of carbonation of the concrete and further its corrosion due to presence of carbon dioxide in standard environment. for this reason, the utilization of hpc is appropriate as prevention against corrosion and concrete damage. due to its excellent durability properties, the long lifespan of constructions is provided, despite exposure to aggressive conditions. the long lifespan and also lower material consumption of hpc structures is the main and most important advantages of this material. however, the primary raw materials which are used for hpc has high environmental impact such as co2 emissions, energy and raw material consumption. the impact on the environment is an important question. nevertheless, new tendencies are heading towards sustainable building and ecological materials. this is the reason why it is appropriate time to think about improving hpc in an environmentally friendly way. silica powder, using as inner filler in hpc, is possible to replace with glass powder, because of similar properties. on one hand, the glass is a nondecomposable material which is necessary to landfilled. however, on the other hand, the material is almost 100% recyclable [3]. recycling is the necessary attitude for the next years. along with recycling it is desirable to minimize the amount of waste [4]. there are several studies [5–8], that examine the possibility of using glass waste as a partial replacement of aggregate in concrete. it is possible to replace coarse aggregate or fine aggregate. the replacement ratio depends on particles size due to possibility of alkalisilica reaction (asr). it is important to notice that on basis of alkali-silica reaction are significant only aggregates with amorphous silica [9]. the utilization of waste glass as replacement of different constituents in concrete mixture has been investigated (aggregate replacement, inert filler, partial cement replacement etc.). the use of waste glass as partial replacement of coarse aggregate for ordinary concrete is possible way in some cases. there are several studies to examine a partial cement and quartz powder (fine fraction) replacement by waste glass powder in ordinary concrete or ultra-high performance concrete (uhpc) [8, 10]. the utilization of waste glass powder as partial replacement in hpc has not been completely explored yet. for this reason, the aim of this investigation was to prove the use of waste glass powder as full replacement of silica powder in hpc. the experiments were focused on a verification of mechanical properties and durability of hpc with waste glass powder. the results were compared with the reference sample containing silica powder. 24 http://dx.doi.org/10.14311/app.2019.21.0024 http://ojs.cvut.cz/ojs/index.php/app vol. 21/2019 glass waste powder utilization in hpc 2. materials and experimental program samples of waste glass powder was collected and examined. there were two types of waste glass – milled waste glass from municipal waste and grinding glass. analysis were made – granulometry and chemical analysis. after that, the grain curves were investigated and compared with the reference sample – silica powder. the same thing was made with chemical composition. after completing analysis of waste glass powder, high performance concrete was made. the basic recipe was optimized in previous research project, cf. table 1. the optimized recipe is used as the reference sample (a). after that, the optimized basic recipe was changed – replacing silica powder with grinding glass (sample b), then with the milled waste glass from municipal waste (sample c). the weight of fine fraction stayed the same. there were two samples (b, c) which were examined and compared with reference sample (a). the differences were evident in the structure of concrete. table 1. basic recipe of hpc – reference sample a. size of prepared samples for tests of mechanical properties was 40 × 40 × 160 mm. there were 54 samples made, 18 from each type. samples 1–15 of each type was ready for tests after 28 days, samples 16–18 was ready for tests after 60 days. it was also meant to compare differences of the mechanical properties between 28-days-old concrete and 60-days-old concrete. 2.1. tests of mechanical properties tests of mechanical properties were made using destructive methods. the tests were made on samples after 28 and 60 days. prepared samples were examined to determine: compressive strength. the tested samples can be cubes, cylinders. in this paper, compressive strength was tested on fragments of samples 40 × 40 × 160 mm. the tested surface ac was 40 × 40 mm. there are certain conditions to keep – about samples (en 12350-1, en 12390-1, en 12390-2 or en 125041), machine (en 12390-4) and procedure. tensile bending strength. tensile bending strength is described in čsn en 12390-5 [11]. the test is realized on block specimens with dimension 40 × 40 × 160 mm. tensile bending strength is defined by three-point press according to čsn en 1015-11 [12]. the test is over after the sample breaks. the highest force registered during the test is used for calculation tensile bending strength. it is always necessary to test at least three samples. freeze-thaw resistance. the tested samples are 40 × 40 × 160 mm. this test was made following čsn 73 1322. cooling is done by air – 4 hours to −20◦c and then 2 hours of warming up to 20◦c. the samples were tested in 25, 50, 75 and 100 cycles. 3. results and discussion the following chapters describe the results and discussion of all tests. 3.1. chemical analysis chemical analysis of samples a, b and c were examined during the spring of 2017. table 2 summarize the results. there are differences such as similarities in chemical composition. after checking amount of sio2 and nao2, cao, it is possible to see difference in number – dozens of percent. chemical composition has an impact to resulting properties. it is necessary to examine potential alkali-silica reaction. the topic of detailed examination of chemical composition could be used in the upcoming research. table 2. results of chemical analysis. 3.2. particle size distribution the size of particle has the main impact to the results. with the growing particle size increases the potential risk of asr. it would be convenient to consider another milling or regrinding of the particles to achieve appropriate size in next research. the comparison of particle size is attached in graph 1. graph 1. comparison of particle size distribution. 25 d. mariaková, t. vlach, t. pavlů acta polytechnica ctu proceedings 3.3. bulk density bulk density of the samples could have the main impact to the results. the porosity in the sample can change the compressive strength or the tensile bending strength values. this research used the same weight amount of waste glass powder to replace silica powder. it would be appropriate to try another attitude in next research – based on bulk density, calculate the appropriate amount of the replacement. graph 2. bulk density of the samples a, b and c – average values from 10 samples. it is possible to see difference between samples b and c compared to sample a, cf. graph 2. however, the difference is negligible – bulk density of sample b is 94, 9%, bulk density of sample c is 94, 3% compared to sample a – considered as 100%. bulk density of the concrete is calculated based on weight and volume. in graph 3, note that the bulk density of the reference sample a is comparable with sample b and sample c. graph 3. bulk density of the concrete after 28 days and 60 days – average values from 3 samples. difference between samples b and c is less than 10% compared to reference sample a in all cases. based on this knowledge, it is possible to compare the samples in following tests. 3.4. tensile bending strength longer hardening time has a positive impact to the results – tensile bending strength increased after 60 days, cf. graph 4. nevertheless, the values between cycles are slightly different (with the increasing number of cycles the values are decreasing), but the impact of freezing cycles is not as big as the impact, for example, particle size or chemical composition. sample b achieved better results than sample c, it is probably caused due to the chemical composition. the values of sample b and c were 68%–88% compared to sample a (considered as 100%). however, it would be appropriate to work on utilization of both samples in hpc and try to improve the results. graph 4. comparison of tensile bending strength – average values from 3 samples. 3.5. compression strength compressive strength values showed positive impact of longer hardening time – the results are better after 60 days (expect sample a – although it is reference sample, in this case it is a negligible value). it would be appropriate to make the other tests based on longer hardening time in next years. the compressive strength values after freeze-thaw resistance tests are satisfying, cf. graph 5, the values of sample b and c range from 70% to 79% compared to sample a (considered as 100%). however, the values are decreasing (sample a, b, c) – it is unusual for hpc. graph 5. comparison of compressive strength – average values from 3 samples. 4. conclusions utilization of waste glass powder in high performance concrete was examined. waste glass powder was used as replacement of silica powder in hpc. mechanical 26 vol. 21/2019 glass waste powder utilization in hpc and physical properties of concrete made with waste glass powder from different sources were investigated. two samples of high performance concrete contain waste glass powder were made, one sample of high performance concrete contains silica powder was made as well – the reference sample. two waste glass powder samples were examined and compared to the reference sample. first, the chemical composition was examined to find out the differences and similarities. there is a possibility to modify chemical composition to improve the results. the particle size distribution was investigated, because the particle size has the main impact to the structure of the concrete and therefore the impact to the results as well. it would be appropriate to consider another milling to reach the right particle size – comparable to silica powder. the bulk density of silica flour, glass powder and concrete was examined as well. the values were slightly different – it should not affect the results of the physical properties. the tests to find out the compressive strength and the tensile bending strength were made. the results were compared to reference sample (contains silica powder). there is a risk of an alkali-silica reaction, because of glass used in the hpc mixture. the freeze-thaw resistance tests confirmed that due to the dense hpc structure, the probability of the alkali-silica reaction was eliminated. this study can be used as a basic for further work. the waste glass powder can be improved in different ways in next years and the utilization in hpc as inner filler is possible. acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs17/010/ohk1/1t/11. references [1] l. myšíčková. vlastnosti, zkoušení a využití uhpc v konstrukcích. brno: vysoké učení technické v brně, fakulta stavební, ústav stavebního zkušebnictví 2011. [2] high performance concrete. http://civil-resources.blogspot.be/2010/06/ high-performance-concrete.html. [3] j. vacková. problematika třídění odpadu u různých věkových kategorií obyvatel. české budějovice 2013. [4] d. vávrová. recyklace skla 2011. [5] a. mohajerani, j. vajna, t. h. h. cheung, et al. practical recycling applications of crushed waste glass in construction materials: a review. construction and building materials 156:443–467, 2017. [6] j. khatib, e. negim, h. sohl, et al. glass powder utilisation in concrete production. european journal of applied sciences 4(4):173–176, 2012. [7] i. b. topcu, m. canbaz. properties of concrete containing waste glass. cement and concrete research 34(2):267–274, 2004. [8] z. z. ismail, e. a. al-hashmi. recycling of waste glass as a partial replacement for fine aggregate in concrete. waste management 29(2):655–659, 2009. [9] r. breitenbücher. alkalicko-křemičitá reakce–důsledky pro cementobetonové kryty. strasse+ autobahn 4, 2006. [10] n. soliman, a. tagnit-hamou. development of ultra-high-performance concrete using glass powder–towards ecofriendly concrete. construction and building materials 125:600–612, 2016. [11] čsn en 12390-5. testing of hardened concrete–part 5: bending strength. czech standardization institute, prague 2007. [12] čsn en 1015-11. zkušební metody malt pro zdivo. 27 http://civil-resources.blogspot.be/2010/06/high-performance-concrete.html. http://civil-resources.blogspot.be/2010/06/high-performance-concrete.html. acta polytechnica ctu proceedings 21:24–27, 2019 1 introduction 2 materials and experimental program 2.1 tests of mechanical properties 3 results and discussion 3.1 chemical analysis 3.2 particle size distribution 3.3 bulk density 3.4 tensile bending strength 3.5 compression strength 4 conclusions acknowledgements references 152 acta polytechnica ctu proceedings 2(1): 152–155, 2015 152 doi: 10.14311/app.2015.02.0152 evolution and outbursts of cataclysmic variables s.-b. qian1,2,3, l.-y. zhu1,2,3, e.-g. zhao1,2, e. fernández lajús4,5, j. zhang1,2, g. shi1,2, z.-t. han1,2 1yunnan observatories, chinese academy of science, p. o. box 110, 650011 kunming, china 2key laboratory for the structure and evolution of celestial objects, chinese academy of sciences, 650011 kunming, china 3university of the chinese academy of sciences, yuquan road 19#, sijingshang block, 100049 beijing, china 4facultad de ciencias astronómicas y geof́ısicas, universidad nacional de la plata, 1900 la plata, buenos aires, argentina 5instituto de astrofisica de la plata (cct la plata conicet/unlp), argentina corresponding author: qsb@ynao.ac.cn abstract mass transfer and accretion are very important to understand the evolution and observational properties of cataclysmic variables (cvs). due to the lack of an accretion disk, eclipsing profiles of polars are the best source to study the character of mass transfer in cvs. by analyzing long-term photometric variations in the eclipsing polar hu aqr, the property of mass transfer and accretion are investigated. the correlation between the brightness state change and the variation of the ingress profile suggests that both the accretion hot spot and the accretion stream are produced instantaneously. the observations clearly show that it is the variation of mass transfer causing the brightness state changes that is a direct evidence of variable mass transfer in a cv. it is shown that it is the local dark-spot activity near the l1 point to cause the change of the mass transfer rather than the activity cycles of the cool secondary star. our results suggest that the evolution of cvs is more complex than that predicted by the standard model and we should consider the effect of variable mass accretion in nova and dwarf nova outbursts. keywords: cataclysmic variables polars photometry individual: hu aqr, dp leo, mn hya, and v2301 oph. 1 introduction cataclysmic variables (cvs) are semi-detached binaries where a low-mass (spectral types of k and m) component star transfers mass to a white dwarf (e.g., warner 1995). according to the standard theory of cvs, the evolution of long-period (p > 3 hours) cvs are driven by angular momentum loss (aml) via magnetic braking (mb), and the ”period gap” (the apparent scarcity of cvs in the period range from 3 to 2 hours) was explained as the cessation of the mb at the period of 3 hours when the secondary component becomes completely convective (e.g., rappaport et al. 1983; spruit & ritter 1983). cvs below the ”period gap” are governed by gravitational radiation. cvs as a mass transferring binary, the aml should keep the secondary filling the critical roche lobe and transferring mass to the white dwarf continuously. in this way, cvs will evolve from a higher mass secondary with longer period to a lower mass secondary with shorter period. mass transfer and accretion in cvs are very important to understand their evolution and observational properties (e.g., the novae and dwarf-novae outbursts). most cvs (e.g., polars and some dwarf novae and novalike cvs) usually show brightness changes in high state, intermediate and low states. the most possible reason to cause these changes is the variation of mass transfer. however, direct evidence is lacking because of the influence of an accretion disk in a normal (non-magnetic) cv. in a polar (magnetic cv), the magnetic field of the white dwarf is very strong to prevent the formation of an accretion disk (e.g., giovannelli & sabau-graziati 2012). the transferred material from the secondary is quickly and directly accreted onto the white dwarf. during the eclipses, light components from the accretion stream and spot on the white dwarf can be isolated. therefore, eclipsing polars can provide invaluable information on the character of mass transfer. the changes of eclipsing profiles directly reflect the variations in the mass-transfer rate. the results will shed light on our understanding of several key astrophysical problems such as the high and low brightness state 152 http://dx.doi.org/10.14311/app.2015.02.0152 evolution and outbursts of cataclysmic variables changes and dwarf nova outbursts. however, only a few polars were monitored photometrically including the brightest polar am her (e.g., hessman et al. 2000; kafka & honeycutt 2003, 2005; wu & kiss 2008), no eclipsing polars were monitored for a long time (e.g., up to several years) before. on the other hand, the eclipse times of eclipsing polars can be determined with a high precision because the accretion stream and the hot spot on the white dwarf can be isolated during the eclipses. very smallamplitude cyclic changes in the observed-calculated (oc) diagram can be detected. therefore, they are also good targets to search for planets and brown dwarfs orbiting cvs by analyzing the light travel time effect. some substellar objects were recently discovered to be orbiting the polars dp leo and hu aqr (qian et al. 2010a, 2011). by using the same method, planets orbiting other evolved binaries were discovered. some examples discovered by our group are those orbiting the detached white-dwarf binaries qs vir and rr cae (qian et al. 2010b, 2012a) and the subdwarf b-type binary ny vir (qian et al. 2012b). 2 targets and observations we have selected ten eclipsing polars, dp leo, ep dra, v2301 oph, ek uma, hu aqr, uz for, v1432 aql, mn hya, v1309 ori, and sdss j015543.40+002807.2, for monitoring photometrically. we monitored those sample stars since 2009 by using the 2.4-m and 1.0m telescopes at yunnan observatories in china and the 2.15-m jorge sahade telescope at complejo astronomico el leoncito (casleo), san juan, argentina. more than 100 eclipse profiles of the selected polars were observed. the eclipse profile for the eclipsing polar, v2301 ophshown in fig. 1. it was obtained on feb. 15, 2012 by using the 2.4-m telescope in lijiang observational station of yunnan observatories. during the observation, no filters were used. among the selected eclipsing polars, we pay more attention to hu aqr. it is one of the brightest eclipsing polars at both optical and x-ray wavelengths. this polar is a total eclipsing binary with an orbital inclination of i = 87◦ and has only one accretion pole (e.g., schwope et al. 2001). all these make it the most ideal target to investigate the properties of mass transfer through the long-term photometry. we have monitored it since 2009 and nearly 70 eclipse profiles were obtained. 3 direct evidence for mass transfer and accretion two of white-light eclipse profiles of hu aqr obtained at high and low brightness states are shown in fig. 2. as shown in the figure, the eclipses start with the limb of the red-dwarf component eclipsing the hot spot (accretion region), and the primary is also eclipsed nearly at the same time. then the accretion stream is the dominant source of the brightness with a small contribution of the secondary, and the distortion in the ingress profile is visible. finally, only the red dwarf is visible and provides a constant contribution at the bottom of the eclipse. 73.440 73.443 73.446 73.449 73.452 73.455 73.458 3.0 2.5 2.0 1.5 1.0 0.5 0.0 -0.5 -1.0 figure 1: white-light photometric observations of v2301 oph obtained on feb. 15, 2012 by using the 2.4m telescope in lijiang observational station of yunnan observatories. solid dots refer to the magnitude differences between v2301 oph and the comparison star, while open circles to those between the comparison and the check stars. -0.10 -0.05 0.00 0.05 0.10 0.15 0.20 4.0 3.5 3.0 2.5 2.0 1.5 1.0 0.5 0.0 figure 2: comparison of two eclipse profiles for hu aqr at high and low brightness states, respectively. the two eclipse profiles were obtained in 2011 by using the 2.4-m telescope in yunnan observatories. the sequence of the egress process is approximately reversed. as displayed in fig. 2, the egress profile of hu aqr is rather stable. therefore, the mid-egress times were used to detect two planetary objects orbiting this binary star by analyzing the light travel time 153 s.-b. qian et al. effect (qian et al. 2011) and the new eclipse times will be used to revise the parameters of the circumbinary planets. as shown in fig. 2, when hu aqr is at a high brightness state, the distortion in the ingress profile is visible. these properties suggest that there is an accretion stream from the secondary to the white-dwarf component and at the same time we see a hot spot on the white dwarf. both the existences of the accretion stream and the hot spot indicate a high mass accretion rate at high brightness state. on the other hand, at a low brightness state, the distortion in the ingress profile disappeared and the eclipse profiles are symmetric. these can be explained as the results that no hot spots on the white dwarf and no accretion streams are seen between the component indicating that no mass transfers and accretions between both components. the brightness variation is correlated with the change of the eclipse profile suggesting that both the hot spot and the accretion stream are produced instantaneously. this is a direct evidence of variable mass transfer and accretion in the cv. two eclipse profiles at high brightness states are plotted in fig. 3. the two eclipse profiles were observed on may 29 and 31,2011 with the 2.15-m jorge sahade telescope in argentina. as displayed in the figure, both the brightness before the eclipse and the ingress profile are variable in two days. other two eclipse profiles obtained in 2010 by using the 1.0-m telescope in yunnan observatories are displayed in fig. 4. it is shown in the figure that hu aqr is on the way from a high state to a low state indicating that the rate of mass transfer is decreasing. 4 discussion and conclusions the observations in previous section give a direct evidence for variable mass transfer and accretion in cvs. the orbital period of hu aqr is about 2.08 hours, just below the cv period gap. non-magnetic cvs with orbital periods in this range are usually dwarf novae (dn). two dwarf novae, ir com and dv uma, have the nearly same orbital period as that of the eclipsing polar hu aqr. they show dn outbursts (a sudden increase of brightness up to 2-7 magnitudes) at usually irregular intervals (nogami et al., 2001). their dn outbursts are explained as the disc instability (cannizzo 1993). however, it is not clear that the outbursts are pure disc phenomena that occur without some contribution of variable mass transfer from the secondary (livio & verbunt, 1988; duschl & livio, 1989). our conclusions suggest that we should consider the contribution of mass transfer in the model of cv outbursts (e.g., cannizzo 1993). the evolution of cvs may be more complex than that predicted by the standard model. -0.20 -0.15 -0.10 -0.05 0.00 0.05 0.10 0.15 0.20 4.5 4.0 3.5 3.0 2.5 2.0 1.5 1.0 oct 22, 2010 figure 3: white-light eclipse profiles observed with the 1.0-m telescope in yunnan observatories. open circles refer to the data obtained on oct. 22, 2010, while solid dots to those obtained on oct. 30. some physical mechanisms, such as irradiationdriven mass transfer and star-spots cause (e.g., wu et al., 1995; livio & pringle 1994), have been proposed to explain the changes of mass transfer in cvs. to solve this question, we analyzed the brightness variation around phase 0.5. it is discovered that the brightness from phase 0.3 to 0.65 usually shows a large dip at low states when the mass transfer rate is very low. two of the light curves obtained at low states are shown in fig. 5 where a large dip is seen around phase 0.5. these brightness dips can be explained by the presence of dark spots near the l1 point on the cool secondary star. this conclusion indicates that it is the coverage of dark spots near l1 point that produce the decrease of the mass transfer rate. 3.52 3.53 3.54 3.55 3.56 3.57 3.58 3.59 3.60 4.5 4.0 3.5 3.0 2.5 2.0 1.5 1.0 0.5 4.51 4.52 4.53 4.54 4.55 4.56 4.57 4.58 4.59 4 3 2 1 figure 4: two light curves of hu aqr observed at low brightness states with the 2.15-m jorge sahade telescope in argentina. it is shown that the brightness around phases 0.5 shows a large dip. the light curve in the left was obtained on nov. 11, 2011, while that in the right was obtained on nov. 12. at the bottom of the eclipse, all of the hot spot, the white dwarf, and the accretion stream are eclipsed by 154 evolution and outbursts of cataclysmic variables the secondary. only the red-dwarf component is visible and provides a constant contribution. therefore, it is the only chance to measure the brightness change of a mass donor in a cv. a cyclic variation in the brightness of the secondary is discovered that can be explained as the dark-spot activity cycle. it is the first time to detect an activity cycle for a fast-rotating mass donor in a cv. our results reveal that the variable mass accretion is not caused by magnetic activity cycles of the fast-rotating red dwarf. acknowledgement this work is supported by chinese natural science foundation through a key project (no.11133007). new ccd photometric observations of those eclipsing polars were obtained with the 2.4-m and 1.0-m telescopes at yunnan observatories (yno) in china and the 2.15-m ”jorge sahade” telescope in argentina. references [1] cannizzo, j. k., 1993, apj 419, 318 [2] duschl, w. j. & livio, m., 1989, a&a 209, 183 [3] giovannelli, f. & sabau-graziati, l., 2012, mmsai 83, 446 [4] hessman, f. v., gänsicke, b. t., mattei, j. a., 2000, a&a 361, 952 [5] kafka, s. & honeycutt, r. k., 2003, aj 125, 2188 [6] kafka, s. & honeycutt, r. k., 2005, aj 130, 742 [7] livio, m. & pringle, j. e., 1994, apj 427, 956 doi:10.1086/174202 [8] livio, m. & verbunt, f., 1988, mnras 232, 1 doi:10.1093/mnras/232.1.1p [9] nogami, d., kato, t., baba, h., novk, r., lockley, j. j., somers, m., 2001, mnras 322, 79 doi:10.1046/j.1365-8711.2001.04114.x [10] qian, s.-b., liao, w.-p., zhu, l.-y., dai, z.-b., 2010a, apj 708, l66 doi:10.1088/2041-8205/708/1/l66 [11] qian, s.-b., liao, w.-p., zhu, l.-y., dai, z.-b., liu, l., he, j.-j., zhao, e.-g., and li, l.-j., 2010b, mnras 401, l34 doi:10.1111/j.1745-3933.2009.00780.x [12] qian s.-b., liu, l., liao, w.-p., li, l.-j., zhu, l.-y., dai, z.-b., he, j.-j., zhao, e.-g., zhang, j., and li, k., 2011, mnras 414, l16 doi:10.1111/j.1745-3933.2011.01045.x [13] qian, s.-b., liu, l., zhu, l.-y., dai, z.-b., fernandez lajus, e., baume, g. l., 2012a, mnras 422, l24 doi:10.1111/j.1745-3933.2012.01228.x [14] qian s.-b., zhu l.-y., dai z.-b., fernández lajús, e., xiang f.-y., and he, j.-j., 2012b, apj 745, l23 doi:10.1088/2041-8205/745/2/l23 [15] rappaport, s., verbunt, f., joss, p. c., 1983, apj 275, [16] schwope, a. d., schwarz, r., sirk, m., howell, s. b., 2001, a&a 375, 419 [17] spruit, h. c., & ritter, h., 1983, a&a 124, 267 [18] warner, b., 1995, cataclysmic variable stars, cambridge univ. press, cambridge [19] wu, k. & kiss, l. l., 2008, a&a 481, 433 [20] wu, k., wickramasinghe, d. t., warner, b., 1995, pasa 12, 60 discussion k. reinsch: which filters did you use for your photometry? could you see a colour dependence in the orbital variation as would be expected for a light curve contribution originating from the secondary? qian shengbang: since we need high time resolution for our data, the exposure time is only a few seconds. therefore, no filters were used during the observations. 155 http://dx.doi.org/10.1086/174202 http://dx.doi.org/10.1093/mnras/232.1.1p http://dx.doi.org/10.1046/j.1365-8711.2001.04114.x http://dx.doi.org/10.1088/2041-8205/708/1/l66 http://dx.doi.org/10.1111/j.1745-3933.2009.00780.x http://dx.doi.org/10.1111/j.1745-3933.2011.01045.x http://dx.doi.org/10.1111/j.1745-3933.2012.01228.x http://dx.doi.org/10.1088/2041-8205/745/2/l23 introduction targets and observations direct evidence for mass transfer and accretion discussion and conclusions acta polytechnica ctu proceedings doi:10.14311/app.2018.18.0072 acta polytechnica ctu proceedings 18:72–76, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app testing of hybrid nickel-polyurethane foams at high strain-rates using hopkinson bar and digital image correlation marcel adornaa, ∗, petr zlámal a, tomáš fílaa, jan faltaa, markus feltenb, michael friesb, anne jungb a czech technical university in prague, faculty of transportation sciences, department of mechanics and materials, konviktská 20, 120 00 prague 1, czech republic b universität des saarlandes, institute of applied mechanics, campus a4.2, saarbrücken, 66123, germany ∗ corresponding author: adorna@fd.cvut.cz abstract. in this paper split hopkinson pressure bar (shpb) was used for dynamic testing of nickel coated polyurethane hybrid foams. the foams were manufactured by electrodeposition of a nickel coating on the standard open-cell polyurethane foam. high strength aluminium alloy bars instrumented with foil strain-gauges were used for dynamic loading of the specimens. experiments were observed using a high-speed camera with frame-rate set to approx. 100-150 kfps. precise synchronisation of the high-speed camera and the strain-gauge record was achieved using a through-beam photoelectric sensor. dynamic equilibrium in the specimen was achieved in all measurements. digital image correlation technique (dic) was used to evaluate in-plane displacements and deformations of the samples. specimens of two different dimensions were tested to investigate the collapse of the foam structure under high-speed loading at the specific strain-rate and strain. keywords: shpb, hybrid foam, dic, strain-gauge measurement. 1. introduction in recent years, we have seen increasing interest in the use of light metal foams in the automotive, aerospace and other applications. this is mainly due to weight reduction, while maintaining load and strength of the construction. another advantage is the ability to achieve better energy absorption capabilities and suitable acoustic damping properties [1]. for design and optimization in the aforementioned applications detailed description of the material properties, particularly under dynamic loads, is necessary. metal foams and similar soft materials in general exhibit a nonstandard deformation behaviour in comparison with conventional materials (e. g. metals) [2], making it difficult to adopt standard measurement techniques for their high strain-rate characteristics and complex deformation behaviour [3–6]. this work is focused on the investigation of the deformation behaviour of nickel coated polyurethane hybrid foam using an innovative shpb device in combination with dic. the presented experimental data describes the course of the foam structure collapse and extends the knowledge of deformation mechanism of hybrid nickel-polyurethane foams, where a coating of nano-crystalline nickel is deposited on cheap open-cell polyurethane foams. 2. materials and methods 2.1. sample preparation polyurethane (pu) foams with a pore size 20 ppi (pores per inch) from jumpinoo, spenge, germany was used. two types of cylindrical specimens with diameter of approx. 20 mm and height approx. 20 mm (long specimens) and approx. 10 mm (short specimens) were cut by hot wire cutting from the foam plates. since pu foams are not electrically conductive, the foams must be rendered electrically conductive by dip coating in a graphite lacquer (crc kontakt chemie, iffezheim, germany) to be able to coat the foams by electrodeposition. the electrodepostion of nanocrystalline nickel (ni) on the carbon coated pu foams was carried out using a commercial ni sulfamate electrolyte with a ni content of 110 g/l (enthone gmbh, langenfeld, germany) at an electrolyte temperature 50 ◦c and ph 3.8. in order to guarantee for a nearly homogeneous coating thickness all over in the foam, the foam was centered as cathode in a double-walled cube filled with ni pellets (a.m.p.e.r.e. gmbh, dietzenbach, germany) acting as sacrificial anode. the electroplating was performed at current density 1.5 ma/cm3 until a theoretical coating thickness of 75µm ni was deposited. for more details, it is referred to the literature [7–9]. the specimen of hybrid foam in the loading device is shown in fig. 1. 2.2. digital image correlation the deformation of the sample during the test was recorded using a high-speed camera. high-speed imaging enables to capture the deformation process of shpb tests in suitable quality (sufficient number of images with small changes between them) for the strain assessment. based on the captured images, changes 72 http://dx.doi.org/10.14311/app.2018.18.0072 http://ojs.cvut.cz/ojs/index.php/app vol. 18/2018 testing of hybrid ni/pu foams using shpb and dic figure 1. ni/pu foam specimen in the shpb experimental setup of the sample structure can be evaluated, using digital image correlation (dic) technique. nowadays, dic technique is a well established method for the assessment of strain in the complex structure using tracking of virtual pattern selected in the sequence of images where the displacements are to be evaluated. each point of the virtual pattern is subsequently tracked through to each pair of images (reference and deformed) in the series of loading states. the computation implementation of the tracking algorithm is based on subsets of pixels which are defined as a square shaped area with defined pixel dimensions formed around the individual tracked point (fig. 2). the size of the subset area has to be suitably chosen with respect to computational cost and traceability. as the specimen is deformed, the reference square pattern is also deformed (combination of translational, rotational and shear deformations) in the image sequence. the position of the deformed subset in the image is calculated based on the extreme of correlation coefficient calculated using the selected criterion. in this work, the algorithm derived from the lucaskanade tracking procedure [10] and implemented in a custom dic tool and the matlab toolkit [11] was used. in this tool, the correlation coefficients were estimated during the two-step procedure, first at pixel level and second at sub-pixel level to determine the values of displacement more accurately. at the pixel level, the sum-squared difference (ssd) criterion was used whereas the lucas-kanade algorithm based on the zero-normalised ssd (znssd) criterion was used at the sub-pixel level. then, strain fields could be calculated from resulting displacements in two dimensions of each point (output of dic algorithm) and mapped to the original structure to describe the inplane deformation behaviour of the sample. 2.3. split hopkinson pressure bar testing of the hybrid foams was performed using a modified kolsky shpb setup. diameter of the measurement bars and the striker bar was 20 mm. all bars were made of high-strength aluminium alloy (enaw-7075). the striker was propelled using a gas-gun with 16 bar maximum pressure. the gas-gun system was composed of the following parts: a steel barrel with the maximum stroke 2500 mm, a high-flow fast figure 2. parameters of the digital image correlation points , parameter m represents the dimension of subset area release solenoid valve (366531, parker, usa), a 20 l air reservoir equipped with pressure gauge and other peripherals (compressor, piping etc.). length of the both incident and transmission bar was 1600 mm. the bars were supported by a set of a low friction polymerliner slide bearings (drylin fjum, igus, germany) mounted in a custom stainless steel housing. the striker bar had length 500 mm. the surfaces of the bars were precisely finished to maintain a tight diameter tolerance for smooth motion in the bearings without notable clearance. the axis of the bars was carefully set to minimize friction and to maintain straightness of the shpb setup. geometrical alignment is very important for the minimization of the undesirable effects (increased friction, bending of the bars etc.) affecting the measurement accuracy. the impact faces of the bars were precisely ground and were set to maintain plan-parallelism between the faces better than 0.05 mm (measured by feeler gauge). an overview of the experimental setup is shown in fig. 3. 2.4. instrumentation the strain wave passing through the incident and transmission bar was measured at two measurement points (mps) using foil strain gauges (3/120 ly61, hbm, germany) with 3 mm active length. the strain gauges were arranged to wheatstone half-bridge circuit to double the output signal in comparison to single gauge and to eliminate the potential influence of bending of the bars. on each measuring bar (1600 mm in length) one mp was set at a distance of 800 mm from each bar end. the foil strain gauges were selected for this application due to their linearity and range of possible strain measurement (up to 50,000µε). the disadvantage of foil gauges is the need to amplify the signal before digitization. this disadvantage can be solved by using semiconductor strain gauges but their nonlinear characteristic and low measurement range (up to 2,500µε) makes them unsuitable for our application. each strain gauge was bonded on the bar 73 m. adorna, p. zlámal, t. fíla et al. acta polytechnica ctu proceedings figure 3. shpb experimental setup overview with a single component low-viscosity cyanoacrylate adhesive (z70, hbm, germany) and cured for at least 12 hours to ensure maximum strength of the bond. to reduce the noise in the measurement chain (maximize to signal-to-noise ratio of the output signal) the powering of the mps was performed using custom battery units with an excitation voltage 4 v. mps signal output were amplified using active differential low noise amplifier (el-lna-2, elsys ag, switzerland) with gain 100 and sampled using high speed 16-bit digitizers (pci-9826h, adlink technology, inc., taiwan) with a maximal 20 mhz sample rate and triggered by signal of through-beam photoelectric sensor (fs/fe 10-rl-ps-e4, sensopart, germany) installed on the barrel of the shpb. the signal from the through-beam photoelectric sensor is also used for the high-speed camera triggering and time synchronization of the strain gauges signal with the captured image sequence. for accurate dic evaluation it is crucial to capture the images in sufficient quality, i.e, properly illuminated and with satisfactory frame rate to ensure a sufficiently small structure changes between the images. thus, the deformation of the sample during the shpb test was observed using a high-speed camera (fastcam sa5, photron, japan) with a selected region of interest (rio) with pixel resolution 256 × 216. the selected roi enabled to reach a frame rate of approximately 105 kfps and 4 pixels resolution per strut thickness of hybrid foam. illumination of the sample surface was performed using a pair of high intensity led lights (constellation 60, veritas, usa). data acquisition was employed using a custom virtual instrument designed in labview environment (national instruments, usa) with tdms binary file output format also readable in the matlab toolkit for further analysis. 3. results 3.1. digital image correlation results two sets of specimen with nominal diameter 20 mm and nominal length 10 mm (short samples) and 20 mm (long samples) were measured using shpb at two different strain-rates. in the paper, the individual strainrates are designated low and high. this designation is relative to the sample. strain-rate achieved during experiments depends not only on striker velocity, but also lenght of tested sample (short samples low strain rate ca. 1500s−1 , high strain rate ca. 5000s−1 , long samples low strain rate ca. 1400s−1, high strain rate ca. 2200s−1). an identical grid of 22 × 14 correlation points was used to cover the area of the specimen and the end parts of the measuring bars on the high-speed camera images. the position of the grid points was tracked using dic and new positions of the grid were identified in the whole sequence of the captured high-speed camera images which were converted from raw format to 8-bit png by a lossless compression algorithm and were not subjected to any additional preprocessing. in all cases, digital image correlation was successful, the mean correlation coefficient was higher than 90 % for all measurements. in fig. 4 median value of the correlation coefficient for three representative experiments: short sample at low strain-rate (test 270), short sample at high strain-rate (test 285), long sample at high strain-rate (test 317) is plotted against strain evaluated using dic. it demonstrates that the trend of all three curves is the same and that the dic results for short samples can be considered reliable up to approx. 0.3 overall strain. in this strain range, the strains evaluated using dic are in very good agreement with the strain-gage measurements. in case of the long samples, the median correlation coefficient exhibits similar trend with dramatic drop of its value for strain higher than 0.2. in strain higher than 0.2 − 0.3, the results provided by dic started to be distorted and unreliable. this phenomena is represented by the drop and the oscillations of the correlation coefficient curves (see green and blue curve in fig. 4) and significant differences in stress-strain curves evaluated using dic and straingauge at strain higher than 0.2 − 0.3 (see fig. 6 -7). evolution of the full-field in-plane strain of the foam structure evaluated using dic is shown in fig. 8 for three selected representative cases. loss of correlation 74 vol. 18/2018 testing of hybrid ni/pu foams using shpb and dic was caused by the densification of the foam structure changing significantly the original structure of the foam. another reason was that diameter of the specimen during compression started to be higher than the diameter of the bars. strain was therefore localized in the central part of the specimen. this distortion caused by the increase of the diameter was even more significant for longer samples (fig. 8 bottom). the dynamic equilirium was achieved in all experiments as the samples were deformed uniformly on both ends in short time after the wave impact on the specimen(see fig. 8 second column of the images). in fig. 8 it is also shown, that hybrid foam deformation occurs locally in areas where the structure is the weakest and collapse of the structure occurs on this plane. figure 4. representative median value of the correlation coefficient plotted against strain for three selected experiments (test 270 short sample at low strain rate, test 285 short sample at high strain rate, test 317 long sample at high strain rate) 3.2. dic and strain-gauge measurement results comparison stress-strain curves with strain derived using dic were evaluated for each tested sample and compared with the strain-gauge curves. a comparison for three measurements (corresponding with full-field dic analysis shown in fig. 8) are shown in fig. 5 to fig. 7 (green lines used for dic curve, orange lines for strain-gauge results). in all cases, stress-strain curves with strain derived using dic are very close to the strain-gauge curves up to 0.3 strain for the short specimens and 0.2 for the long specimens. differences of the dic curves to strain-gauge curves after 20-30 % deformation are caused by the fact that sample size exceeds the measuring bars diameter, resulting in a loss of correlation points in parts outside the monitored area. moreover, significant deformation of the sample also leads to a loss of correlation points due to the collapse of the structure and shift of individual struts out of high-speed camera focus. 4. conclusion shpb experimental device was successfully used for the dynamic compression of the ni/pu hybrid foam. in-plane strain was successfully evaluated from the figure 5. comparison of stress-strain diagrams with strain evaluated using strain-gage and dic, experiment no. 270 figure 6. comparison of stress-strain diagrams with strain evaluated using strain-gage and dic, experiment no. 285 figure 7. comparison of stress-strain diagrams with strain evaluated using strain-gage and dic, experiment no. 317 high-speed camera images using digital image correlation technique. dic results were found to be in a very good agreement with strain-gauge measurements up to 30 % of deformation for the short samples and 20 % of deformation for the long samples. dynamic equilibrium was achieved in all measurements as the specimens were deformed uniformly on both ends in short time after the wave impact on the specimen. it was shown that both shpb and dic are suitable tools for advanced characterization of the hybrid foam under impact loading. 75 m. adorna, p. zlámal, t. fíla et al. acta polytechnica ctu proceedings figure 8. full-field dic analysis: in-plane longitudinal strain top row: experiment no. 270, vs = 22.1 ms−1 (low strain-rate, short sample), d = 22.4 mm, l = 10.1 mm middle row: experiment no. 285, vs = 38.1 ms−1 (high strain-rate, short sample), d = 21.9 mm, l = 7.5 mm bottom row: experiment no. 317, vs = 43.7 ms−1 (high strain-rate, long sample), d = 21.8 mm, l = 18.6 mm vs [ms−1] striker velocity, d [mm] sample diameter, l [mm] sample length acknowledgements the research was supported by the czech science foundation (project no. 15-15480s) and the internal grants of the czech technical university in prague (projects no. sgs17/148/ohk2/2t/16 and no. sgs18/153/ohk2/2t/16)). all the financial support is gratefully acknowledged. references [1] t. fila, p. zlamal, o. jirousek, et al. impact testing of polymer-filled auxetics using split hopkinson pressure bar. advanced engineering materials 19(10), 2017. doi:10.1002/adem.201700076. [2] g. sunny, f. yuan, v. prakash, j. lewandowski. design of inserts for split-hopkinson pressure bar testing of low strain-to-failure materials. experimental mechanics 49(4):479–490, 2009. doi:10.1007/s11340-008-9145-1. [3] p. wang, s. xu, z. li, et al. experimental investigation on the strain-rate effect and inertia effect of closed-cell aluminum foam subjected to dynamic loading. materials science and engineering a 620:253–261, 2014. doi:10.1016/j.msea.2014.10.026. [4] a. jung, m. larcher, o. jirousek, et al. strain-rate dependence for ni/al hybrid foams. vol. 94. 2015. doi:10.1051/epjconf/20159404030. [5] k. dannemann, j. lankford jr. high strain rate compression of closed-cell aluminum foams. materials science and engineering a 293(1):157–164, 2000. doi:10.1016/s0921-5093(00)01219-3. [6] v. deshpande, n. fleck. high strain rate compressive behaviour of aluminum alloy foams. international journal of impact engineering 24(3):277–298, 2000. doi:10.1016/s0734-743x(99)00153-0. [7] a. jung, h. natter, s. diebels, et al. nanonickel coated aluminum foam for enhanced impact energy absorption. advanced engineering materials 13(1–2):23–28, 2011. doi:10.1002/adem.201000190. [8] a. jung, k. m. r., e. lach, et al. hybrid metal foams: mechanical testing and determination of mass flow limitations during electroplating. international journal of material science 2(4):97–107, 2012. doi:10.1002/adem.201000190. [9] a. jung, s. diebels. synthesis and mechanical properties of novel ni/pu hybrid foams: a new economic composite material for energy absorbers. advanced engineering materials 18(4):532–541, 2015. doi:10.1002/adem.201500405. [10] b. d. lucas, t. kanade. an iterative image registration technique with an application to stereo vision. in proceedings of imaging understanding workshop, pp. 674–679. 1981. [11] i. jandejsek, j. valach, d. vavrik. optimization and calibration of digital image correlation method. pp. 131–136. 2010. 76 http://dx.doi.org/10.1002/adem.201700076 http://dx.doi.org/10.1007/s11340-008-9145-1 http://dx.doi.org/10.1016/j.msea.2014.10.026 http://dx.doi.org/10.1051/epjconf/20159404030 http://dx.doi.org/10.1016/s0921-5093(00)01219-3 http://dx.doi.org/10.1016/s0734-743x(99)00153-0 http://dx.doi.org/10.1002/adem.201000190 http://dx.doi.org/10.1002/adem.201000190 http://dx.doi.org/10.1002/adem.201500405 acta polytechnica ctu proceedings 18:72–76, 2018 1 introduction 2 materials and methods 2.1 sample preparation 2.2 digital image correlation 2.3 split hopkinson pressure bar 2.4 instrumentation 3 results 3.1 digital image correlation results 3.2 dic and strain-gauge measurement results comparison 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.10.0038 acta polytechnica ctu proceedings 10:38–42, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app temperature of soil monitored at experimental sites in perm region (russia) andrey ponomaryov∗, aleksandr zakharov perm national research polytechnic university, kuibysheva street 109, perm, russia ∗ corresponding author: spstf@pstu.ru abstract. the paper presents the monitoring results for soil temperatures in the geological and climatic conditions of the city of perm (russia). the monitoring results are presented for two experimental sites. site 1 is characterized by dense urban development and related to type i soil conditions in perm (the left-bank area of perm). site 2 is located in a low-urbanised region of the city and is related to type ii soil conditions in perm (the right-bank area of perm). keywords: temperature of soil, monitoring site, energy-efficient foundations. 1. introduction the department of construction technology and geotechnics of perm national research polytechnic university has conducted studies of the temperature fields on ground bases in the city of perm since 2009. research is being carried out to quantify the thermal behavior of soil for the main types of geotechnical condition. quantitative assessment of subgrade temperature is one of the main factors allowing introduction of a process based on the use of the soil thermal energy. in particular, these are the energy-efficient foundations and underground structures of buildings. as a rule, technologies based on the use of soil thermal energy are employed for the heating and cooling of buildings, at the very least to provide electricity and hot water. these technologies have been widely used in european and neighboring countries [1–5]. their introduction in russia at this point is not yet much in evidence. 2. geological section of experimental sites research of temperature fields on ground bases are being carried out for two basic types of geotechnical conditions specific to the left-bank and right-bank areas of perm city. studies are being conducted at two sites: • site 1 the monitoring system is installed at the civil engineering faculty pnrpu in the sverdlovsky district of perm (left bank of the city). monitoring has been carried out continuously from december 2008 to the present time. site 1 is characterized by dense urban development. the development age is more than 50 years. the distance from the observation well to the nearest building is about 3 m. taking the results of a geological engineering survey, experimental site 1 is comprised of quaternary alluvialdealluvial clayey soil of a total thickness of 11.6 m, with up to 60-70 % of pebbles at the bottom. quaternary soils are overlaid by filled-up ground of thickness 6.0 m. filled-up ground is represented by clay soil of semi-solid to plastic consistency at the base with an admixture of 60-70 % of construction waste (rubble, broken bricks, glass, wood). the bedrock is argillites uncovered at a depth of 17.6 m. quaternary alluvial-dealluvial deposits are mainly represented by clay of solid to semi-solid consistency underlying the pebble with loamy sand of solid consistency (the content of gravel and pebbles being 60-70 %). the geological section is shown in fig. 1a. • site 2 the monitoring system is installed on the territory of the university campus in the leninsky district of perm (the right bank of the city). monitoring has been carried out since august 2015. currently monitoring data has been obtained up to january 2016. site 2 is located on the undeveloped area of the city. the distance to the nearest building is more than 30 m. according to the results of archival research of the surrounding area, site 1 is comprised of quaternary alluvial sand and clay of a total thickness of 15 m, with 25 % of gravel at the bottom. quaternary alluvial deposits are represented by fine sand, clay and clay soils of plastic to liquid consistency. the gravel inclusion was marked from a depth of 1.0 m along the full thickness of the alluvial soils with an increase to the base up to 25 %. according to the results of archive research the bedrock was uncovered at a depth of 17.0 m. the geological section is shown in fig. 1b. according to the results of geotechnical studies sites 38 http://dx.doi.org/10.14311/app.2017.10.0038 http://ojs.cvut.cz/ojs/index.php/app vol. 10/2017 temperature of soil monitored at experimental sites figure 1. geological section of experimental sites 1 and 2 are assigned to the i and ii types of geotechnical conditions respectively, specific to perm [6]. 3. monitoring results of the temperature fields studies of the temperature field distribution in the soil were carried out by installing a resistive temperature transducer (temperature sensors). installation of this resistive temperature transducer was made in a predrilled hole with a protective casing. to gather the data rtm 59 loggers were used for measuring, continuous recording and monitoring of temperature and other non-electrical values (frequency, pressure, flow, level, etc.), converted into electrical signals of strength, dc voltage and dc resistance [7]. the thickness of the observed soil mass on the site 1 is 19 m. the temperature sensors were installed at intervals of 1 m. monitoring of temperature fields has been held for more than 5 years on site 1. the summarized monitoring results (average monthly temperatures for the period from january 2009 to january 2016) are shown in fig. 2. as given in the diagram the temperature of the soil mass from a depth of 8-9 m is practically independent of the seasonal outdoor temperature fluctuations. the temperature is about 12 °c, falling to 10 °c to the depth of 19 m. to analyze the temperature fluctuations for several annual cycles the average monthly temperatures in september over the period 2009-2015 is shown in fig. 3. as given in the diagram the temperature of the soil mass from a depth of 8-9 m is constant (the difference is not more than 1 °c) during 6 years of monitoring. soil temperatures in september 2015 are explained by abnormally low outside temperatures in summer 2015. the thickness of the observed soil mass on the site 2 is 37 m. the temperature sensors were installed 39 andrey ponomaryov, aleksandr zakharov acta polytechnica ctu proceedings figure 2. site 1. diagram of average monthly temperatures of soil mass according the monitoring results from 2009 to 2016 figure 3. site 1. diagram of average monthly temperatures of soil mass in september for the 2009-2015 at intervals of 2 m, to a depth of 3 m with intervals equal to 0.5 m. monitoring of the temperature fields on site 2 has been carried out since august 2015. currently monitoring data of temperature distribution in the soil mass has been obtained up to january 2016. monitoring results (average monthly temperatures) are shown in fig. 4. analysis of monitoring results on site 2 displays the temperature of soil mass from a depth of 9 m as constant and equal to 6-7 °c. for further analysis the average monthly temperatures of soil mass for september 2015 are shown in table 1 and fig. 5 for both sites. 4. conclusions based on the monitoring of the temperature regime of soils, conclusions are drawn: (1.) the temperature of the ground for the second site is + 6-7 °c and does not change over a depth of 9 m. (2.) the temperature of the ground of the first site is + 10-12 °c and does not change during the year from a depth of 9 m. with the depth reading, the temperature decreases. (3.) on the temperature regime of soils, the density of development of the territory under study has a significant influence. the higher temperatures of the ground massif of the first site are due to the proximity of the building of the faculty of civil engineering with a heated basement. 40 vol. 10/2017 temperature of soil monitored at experimental sites figure 4. site 2. diagram of average monthly temperatures of soil mass according the monitoring results from august 2015 to january 2016 figure 5. diagram of average monthly temperatures in september for sites 1 and 2 table 1. average monthly temperatures of soil mass in september for sites 1 and 2 41 andrey ponomaryov, aleksandr zakharov acta polytechnica ctu proceedings references [1] h. brandl, d. adam, r. markiewicz. ground-sourced energy wells for heating and cooling of buildings. acta geotechnica slovenica 3(1):5–27, 2006. [2] r. katzenbach, d. adam, t. waberseck. on the geotechnical conditions of construction in the city of perm. problems of soil mechanics and foundation engineering in difficult ground conditions. in proc. of the international scientific conference dedicated to the 50th anniversary of bashniistroy 3-5 october. ufa: bashniistroy, 2006. [3] v. g. shapoval, b. v. morklyanyk. the temperature field in the soil base of heat pumps: monograph. dnepropetrovsk: porohy, 2011. [4] a. b. ponomaryov, a. v. zakharov. use of geothermal energy of soils for heating and air-conditioning of buildings. bulletin of volgograd state university of architecture and civil engineering 17(36):119–122, 2010. [5] a. zakharov. application of geothermal energy of the ground for heating of buildings in the climatic and engineering geological conditions of the perm region. bulletin of civil engineers 2 (23) 2(23):85–89, 2010. [6] r. katzenbach, t. waberseck, d. adam, a. ennigkeit. innovationen bei der nutzung geothermischer energie durch erdberührte bauwerke, wie z.b. pfahlgründungen mittels energiepfählen. geothermie-symposium "erdwärme energieträger der zukunft", 25.11.2002 bremerhaven / bremerhavener energiemanagement-agentur. 2002. [7] a. b. ponomaryov, a. v. zakharov. analysis of the interaction of energy foundations in the geological and climatic conditions of the city of perm. pnrpu bulletin 4:24–33, 2011. 42 acta polytechnica ctu proceedings 10:39–43, 2017 1 introduction 2 geological section of experimental sites 3 monitoring results of the temperature fields 4 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0074 acta polytechnica ctu proceedings 12:74–78, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app correlation analysis of vehicle frontal impact parameters josef mík∗, jana kadlecová department of vehicle technology, čvut v praze, konviktská 20, 12000 prague ∗ corresponding author: mikjosef@fd.cvut.cz abstract. the article considers a possible improvement of road vehicle safety by using ecall – a system which initiates an emergency call in case of traffic accident. a possible way of better description of a frontal impact accident of a vehicle is examined and enriched by the information from the onboard e-call unit. in this article, we analyze results of frontal crash tests with different types of barriers and overlapping area and look for the correlation between the individual vehicle and collision parameters in order to provide a better description of the severity of the accident by the ecall system. the relation among the selected parameters is described using the correlation analysis. keywords: crash test, ecall, cotrrelation. 1. introduction in a long-term timeframe, there is a noticeable decrease in the amount of road accidents, casualties and fatal victims of road accidents in europe. it is mostly achieved thanks to newly accepted legislation measures increasing road safety and minimizing consequences of road accidents. one of these measures is a mandatory installation of the ecall system – an emergency call system in case of a road accident. the system is to be launched and installed on all newly produced vehicles in april 2018 [1]. based on the information obtained from passive safety systems installed in the vehicle and additional ecall system sensors, the ecall system automatically detects a road accident and consequently sends a minimum set of data (msd) message to the emergency service and at the same time establishes a voice communication with 112 service. the operator of the emergency service decides on the measures and safety services dispatched to the accident site based on the information obtained in the msd message and from oral communication with passengers of the crashed vehicle. the basic structure of the msd message contains the following information about the vehicle and the accident: • version of msd format; • identification of msd; • ecall activation type; • vin code for vehicle (iso 3779)[2]; • fuel type; • vehicle direction; • recent vehicle location n1, n2; • number of passengers (locked safety belts). a set of additional information, which can be used for more accurate description of consequences of accident, is described in this article. this was concluded based on results of series of crash tests. 2. experiment description a background for the analysis is data obtained from a series of front impact collision of vehicles category m1. results of these crash tests are shown in table 1. calculated resulting injury criteria coefficient hic and 3 ms head criteria are represented in table 2. as it can be seen from table 1 and 2, the constellations of test vehicles contain vehicles equipped with different types of passive safety systems. activation of the frontal airbag was a monitored parameter. as for the collision partner, a solid barrier with a 100 the following parameters were selected for correlation analysis: • overlapping of the frontal area – indicates overlapping of the frontal area of the test vehicle with the collision partner in percent; • collision velocity (v0) – calculated based on the video of the collision from high-speed cameras; • the weight of the vehicle – calculated based on measurements performed on separate wheels with plate weights; • kinetic energy of the vehicle(ek) – is calculated based on the vehicle weight, and collision velocity (equation 1) ek = 1 2 ∗m∗v2, (1) where: ek – kinetic energy of the vehicle right before the collision [j], m – weight of the vehicle [kg], v – collision velocity [m*s−2]; • maximal acceleration of the vehicle body – estimated from the acceleration of the vehicle body during the collision in x direction; 74 http://dx.doi.org/10.14311/app.2017.12.0074 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 correlation analysis of vehicle frontal impact parameters test vehicle collision vehicle characteristicspartner overlap [%] front airbag v0 [km*h−1] m [kg] ek [kj] amax [m*s−2] deformation [mm] škoda fabia rigid barrier 100 yes 50,77 1208 120,13 620 447 škoda favorit rigid barrier 100 no 47,3 875 75,53 459,8 508 škoda rapid rigid barrier 100 yes 45 1235 96,48 378,1 416 škoda rapid rigid barrier 100 no 14,92 1122 9,64 115,6 127 renault 5 gtd rigid barrier 40 no 52,2 938,6 98,67 345 700 škoda roomster deform. barrier 40 yes 56,18 1542 187,76 no no škoda fabia škoda octavia 33 yes 50 1157 111,59 176 815 table 1. results of crash tests – characteristic of test vehicles. test vehicle collision partner driver manikin other manikinamax [m*s−2] hic [-] location in the vehicle amax [m*s−2] hic [-] škoda fabia solid barrier 510 302,94 passenger 2277 4139 škoda favorit solid barrier 2490 2205,6 child p3 708,8 706 škoda rapid solid barrier 460 180,42 ne ne no škoda rapid solid barrier no ne ne ne no renault 5 gtd solid barrier no no child p3 856,6 612 škoda roomster deform. barrier no 359,25 passenger ne 358,64 škoda fabia škoda octavia no no no no no table 2. results of crash tests – hic criteria of test manikins. • resulting deformation of the vehicle (ξ) – estimated from visual documentation of the test vehicle after the accident (photos and video from the high-speed cameras), based on comparison with photo documentation of the test vehicle before the collision; • maximal acceleration of the manikin driver’s head (amax) – estimated from the acceleration acting on the driver manikin’s head during the impact. the resulting acceleration curve was calculated from the accelerations of the manikin head in x, y and z axis based on the equation 2. the overall progress of result acceleration was then filtered with standard frequency filter cfc1000. acelk = √ a2x + a2y + a2z, (2) where: acelk – resulting acceleration [m*s−2], ax,ay,az – acceleration of the manikin head measured in x, y, z [m*s−2]; • hic criteria value for the driver manikin – head injury criteria was calculated from the result acceleration curve based on the equation 3. in case there was no contact of the manikin head with the vehicle interior the time interval of 36 ms was selected for time t1 to t2 (estimated from the speed camera recordings), in casethere was a collision, a 15 ms time interval t1 to t2 was selected. hic criteria value is then the maximum value in the selected interval for the result acceleration hic = [ 1 t1 − t2 ∗ ∫ t2 t1 adt ]2,5 ∗ (t1 − t2), (3) where: hic – value of head injury criteria [-], a – result acceleration [g], t1 – start of the time interval for hic calculation, t2 – end of the time interval for hic calculation. results of crash tests are summarized in tables 1 and 2. 3. correlation analysis of crash test results analysis of the crash tests results was made based on correlation analysis between selected parameters presented in table 1 and 2, considering if it is suitable to use them for estimation of the severity of the accident by ecall system. the relationship between parameters is expressed by a regression function using the least squares method (equation 4) 75 josef mík, jana kadlecová acta polytechnica ctu proceedings yi − ŷi yi −ayx + byx ∗xi ... yi − ŷi · yi −ayx − byx ∗xi ... f ( ayx,byx ) = 1 n ∑n i=1 (4) where: ŷi – aligned value (theoretical), yi – emperical values, ayx,byx – equation coefficients. the reliability of regression estimation was determined by the correlation coefficient ryx. correlation dependency is considered to be poor tight if the value of correlation coefficient is lower than 0.3. in case the value is higher than 0.8, then correlation dependency is considered to be very tight. if the value lies in the interval from 0.3 to 0.8 the correlation dependency is considered to be mild tight. 4. dependency of the vehicle body deformation from collision velocity the dependency of the deformation of the vehicle from collision velocity is represented in figure 1. from all the crash tests, only collisions with a rigid barrier with 100% overlapped were used for correlation analysis. these values are shown in the figure as blue dots with a line regression function. it is possible to calculate correlation coefficient from the coefficient of determination r2 as it is shown in the equation 5. r = √ r2 = √ 0, 9433 = 0, 9712 (5) the resulting value of the regression coefficient is r = 0.9712 so we can consider correlation dependency as very tight. but a small amount of input data has to be considered. in figure 1 there are also shown other values which were not used for the regression analysis because of a different type of collision parameters. these are the data obtained from crash tests with the vehicle renault 5 gdi, which was collided with a rigid barrier with 40% overlap, and a skoda fabia sedan, which was collided with another vehicle skoda octavia with 33% overlap. it is visible from figure 1 that the value of the resulting deformation is increasing with decreasing overlap of the collision area. 5. the dependency of maximum acceleration on the kinetic energy of the vehicle the next parameter that we consider in this research is the dependency between the value of maximum acceleration, which impacts the vehicle body during a crash, and the value of the vehicle’s kinetic energy, which is transformed into the vehicle’s body deformation during the collision. the values for individual vehicles are incorporated in the graph in figure 2. as in the case of the dependency between vehicle body deformation and collision velocity, the regressive function describes only vehicles that suffered a frontal collision with a rigid barrier with 100% overlap. the values that are applicable to personal vehicles can also be found in the graph. to describe the dependency, an exponential regression function was chosen as it has a higher value for correlation coefficient than a linear function. the graph describing the equation of the regression function can be found in figure 2, as well as the determination coefficient r2. equation 6 establishes the value of the correlation coefficient through a determination coefficient. r = √ r2 = √ 0, 9103 = 0, 954 (6) according to the resulting value of the correlation coefficient r = 0,954 it is possible to see the correlation dependency as very high. however, it is necessary to take into consideration the small number of input values on the basis of which the function was solved. the graph in figure 2 also contains the values of acceleration for the renault 5 gtd (40% overlap) and the škoda fabia sedan (33% overlap) with the lesser percentage of overlap. the graph demonstrates that the maximum acceleration influencing these vehicles with an equal value of the consumed kinetic energy in the case of vehicles with a 100% overlap is smaller. in the case of the škoda fabia sedan, however, part of the kinetic energy was given to another vehicle, škoda octavia, that was moved by 9,28 meters from its original position before the crash, as well as accelerated. for this reason, the tested vehicle was slowed down, which does not happen in case of frontal collision with a rigid barrier. this parameter, therefore, should not be considered. 6. the dependencies of the hic criteria on the speed of the vehicle in this chapter, the correlation of the possibility of the driver’s head injury and the speed of the vehicle is considered in order to study the severity of injuries passengers of the car can suffer during a traffic accident. as it is clear from the results of the car collision that an airbag can have a serious impact on the resulting value of the hic criteria, the correlation assumes that the airbag was used. the graph in figure 3 presents values which also include the values established for the passenger mannequin. the linear regression function presented in the graph is applicable to a restraint system of a vehicle that has an airbag. the quantity of the correlation coefficient is calculated by equation 7. r = √ r2 = √ 0, 9024 = 0, 9499 (7) 76 vol. 12/2017 correlation analysis of vehicle frontal impact parameters figure 1. dependency of the deformation of the vehicle from collision velocity. figure 2. dependency of the deformation of the vehicle from collision velocity. 77 josef mík, jana kadlecová acta polytechnica ctu proceedings figure 3. dependency of the deformation of the vehicle from collision velocity. according to the resulting value of the correlation coefficient r = 0,9499 it is possible to see the correlation dependency as very high. however, it is necessary to take into consideration the small number of input values on the basis of which the function was solved. to establish the regression function we have used only 3 input hic values. in the case where we need to describe the seriousness of the injuries suffered by the passengers in a traffic accident, it is better to use the ais method instead of the resulting value of the criterion of the injuries. the ais method describes the seriousness of the injuries suffered by individual parts of the human body using a number in the 0 – 6 interval. this possibility is considered in the research described by [3] in the bibliography. 7. conclusion the problem of the estimation of the consequences arising from a vehicle collision was solved on the basis of establishing a dependency between the chosen vehicle parameters. in the chosen pairs of the parameters one quantity was possible to measure with ecall while another quantity was to be estimated. the dependency between the chosen parameters was described with a regression function and its accuracy was verified on the basis of the value of the correlation coefficient. in all three cases the accuracy of the quantity’s dependence of the correlation coefficient rxy was verified. in all three cases, this coefficient came out to be higher than 0,8 and therefore we can establish the dependency between the chosen quantities to be very high. however, in considering the outcome we also have to bear in mind the low number of quantities on the basis of which the correlation analysis was carried out. this fact considerably undermines the reliability of the results. to strengthen the reliability of the results it is necessary to carry out a correlation on the basis of a higher number of crash tests with different impact velocities. another way to acquire data for statistical analysis is to create a fem or multi-body model and to carry out a simulation of collision impact with simulation software. this would also allow for the repetition of the scenario with the same model and an easy change in input parameters. references [1] e. 2015/758. požadavky na schválení typu pro zavedení palubního systému ecall využívajícího linku tísňového volání 112. [2] i. 3779:2009. road vehicles – vehicle identification number (vin) – content and structure. [3] h. kovandová, h. krejčí, j. kovanda. crash tests of vehicles – selected experiments and analysis. in: transactions on transport sciences, vol. 6, no. 4/2013, issn: 1802-9876, pg. 193-202. 78 acta polytechnica ctu proceedings 12:74–78, 2017 1 introduction 2 experiment description 3 correlation analysis of crash test results 4 dependency of the vehicle body deformation from collision velocity 5 the dependency of maximum acceleration on the kinetic energy of the vehicle 6 the dependencies of the hic criteria on the speed of the vehicle 7 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2018.14.0042 acta polytechnica ctu proceedings 14:42–48, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app imitating neutron irradiation damage using a combination of ion beams václav šísla, ∗, martin ševečeka, b a department of nuclear reactors, faculty of nuclear sciences and physical engineering, czech technical university in prague, v holešovičkách 2, prague 8, czech republic b department of nuclear science and engineering, massachusetts institute of technology, 77 massachusetts ave, cambridge, ma, usa ∗ corresponding author: sislvacl@fjfi.cvut.cz abstract. there is a strong motivation for using ion beams to imitate neutron irradiation damage, mainly in order to reduce costs and time demands linked to neutron irradiation experiments. the long-term goal of the authors is to create an ion irradiation methodology, which could be employed in the development process of innovative nuclear fuel materials. this methodology will be based on combining of a set of ion beams in such a way that the final distribution of irradiation-induced defects in the material structure is similar to the one which would have been introduced by neutrons in a nuclear reactor. the first part of the methodology is represented by an optimization tool described here. the tool uses a third party monte carlo code srim to simulate ion transport in a target and to determine the distribution of radiation damage. subsequently, a custom genetic optimization algorithm is applied to a set of damage distribution profiles to find their optimal combination. keywords: accident tolerant fuel, fuel cladding, radiation damage, srim, defects distribution optimization, genetic algorithms. 1. introduction development of innovative materials for nuclear reactor applications (such as new cladding materials) is a complex multidisciplinary process. one of the criteria to be fulfilled is to prove resistance of the material against neutron irradiation and to quantify material characteristics that degrade with radiation. a traditional approach of proving irradiation tolerance is through irradiation in a research reactor and subsequent post-irradiation experimental testing (pie). however, using a research reactor for irradiation purposes has multiple disadvantages, of which the most significant ones are time demands and costs. the time demands are the result of two factors. first, the irradiation time itself is long, as the damage rate in a nuclear reactor is usually low, thus it takes a long time to accumulate the desired damage (often several months or even years). second, since the samples become radioactive, further time is needed for decaying of the samples before subsequent testing. the fact that samples become radioactive also results in requirements for special handling procedures and facilities certified for such purposes. another drawback of neutron irradiation is the need for designing of special capsules, in which the samples are placed during irradiation or need to handle the material in hot cells. this further increases the costs and complexity of the experiments. there are also advanced nuclear reactors under development around the world. the high dose reactor figure 1. comparison of approximated requirements for temperature and dpa for advanced reactor concepts and annual damage rates produced by current research and fast reactors. [1] concepts as molten salt reactors or traveling wave reactor are very challenging from the perspective of radiation damage of structural components that must survive radiation damage up to 250 dpa or higher. [1] this makes the development of new materials very complicated and leads to excessive prolongation of the development of advanced reactors. the requirements for temperature and level of radiation damage to be reached for different advanced reactors are summarized in figure 1. taking into account all the problems linked to neutron irradiation in research reactors, there is a strong motivation for using ion beams. the main advantages 42 http://dx.doi.org/10.14311/app.2018.14.0042 http://ojs.cvut.cz/ojs/index.php/app vol. 14/2018 simulation of irradiation processes of ion irradiation can be summarized as: • sources of ion beams (accelerators) are widelyspread. • dose rates are much higher compared to neutron irradiation, thus exposure times are as low as hours or tens of hours. • there is no need for special capsule design. • there is only a low or none residual activity of irradiated samples. • experiment conditions can be well controlled easily adjustable ion energy, dose rate, and sample temperature. • damage cascades produced mainly by primary knock-on atom (pka) in the case of irradiation by heavy ions of typical energies are similar to those created by a pka knocked by energetic neutrons. there are, however, differences between the effects of neutron and heavy ion irradiation. these will be discussed in the following section. the goal of current research is to achieve as comparable irradiation damage profile as possible resulting in imitation of irradiation damage by heavy ions. 2. differences in effects of ion and neutron irradiation in spite of the fact that ion irradiation offers significant advantages compared to neutron irradiation, there is a fundamental question to be answered: are the effects of ion irradiation comparable to those created by neutrons? according to was [2], even though there are significant differences between the direct effects of neutron and ion irradiation, it is possible, to a certain degree, imitate the effects of neutron irradiation using ion beams. an important fact is that it is the final state of the material that determines the equivalence, not the path taken. the main differences between effects of ion and neutron irradiation are [1–3]: • particle energy spectrum: neutrons in a reactor have a wide energy spectrum ranging between several orders of magnitude from thermal energies (0.025 ev) up to fast neutrons (mevs). on the other hand, heavy ions produced by an accelerator have very narrow range of energy (usually considered as monoenergetic beams). • penetration depth and damage distribution: neutron’s mean free path in structural materials of nuclear reactors is in order of centimeters, thus they can penetrate deep in a target. a large mean free path also means that the spatial damage distribution created by neutrons is homogeneous over many millimeters of the penetration depth. ions, on the other hand, due to their electric charge, lose energy much faster when traveling through a material figure 2. damage distributions caused by various energetic particles. [2] and their penetration depths are orders of magnitude shorter. they usually penetrate between 0.1 and 100 µm depending on the particle type and its energy, see figure 2. • damage rate: as the cross-section for ion-atom interaction is much higher compared to neutron-nuclear interaction, the damage rate relative to the particle fluence is orders of magnitude higher for ion irradiation than for neutron irradiation. • nature of induced defects: neutrons and heavy ions produce large damage clusters of comparable size and nature. light ions and protons produce smaller clusters or isolated point defects and electrons produce widely separated self-standing point defects. this behavior is more complex due to diffusion related effects that are closely linked to the temperature of the target. in the presented study, we decided to tackle the issue of different damage distributions resulting from neutron and ion irradiation. our goal was to combine a set of several ion beams of different energies and/or ion types. assuming that the damage profiles created by individual beams can be summed-up independently, it is then possible to find a particular optimized combination leading to a relatively flat damage profile similar to the one, which would have been created by neutrons. the proposed optimization method is described in the following section, an example of a typical damage profile created by a single ion beam (6.0 mev, ni) and an optimized damage profile, as we defined it, are shown in figure 3. in the past, there were multiple studies investigating a possibility to combine irradiation by multiple ion beams to mimic neutron irradiation [3–5]. however, the main aim was not to obtain a uniform damage 43 václav šísl, martin ševeček acta polytechnica ctu proceedings 0 1 2 3 4 5 6 7 8 9 0 5 × 10−16 1 × 10−15 1.5 × 10−15 2 × 10−15 dp a pe r flu en ce [d pa ·c m 2 / io n] target depth [µm] ti2alc + zircaloy-4, dpa per fluence ni, 6.0 mev cladcoat 0 1 2 3 4 5 6 7 8 9 0 5 × 10−16 1 × 10−15 1.5 × 10−15 2 × 10−15 dp a pe r flu en ce [d pa ·c m 2 / io n] target depth [µm] ti2alc + zircaloy-4, dpa per fluence ideal damage distribution cladcoat figure 3. damage profile created by 6 mev nickel ions irradiating 3 µm thick ti2alc coating layer deposited on zircaloy-4 slab (top) and an optimal damage profile as defined by the authors (bottom). distribution, but rather to address neutron-induced transmutation by injecting he while irradiating by another type of ions to obtain the desired level of damage. also, a procedure combining neutron irradiation for obtaining radiation damage and he irradiation for injecting bubbles is being used by the community. [6] the possibility of obtaining a layer1 of evenly irradiated material might be especially suitable for testing accident tolerant fuel (atf) multi-component cladding materials. during such tests, it is needed to evaluate properties under irradiation of a thin coating layer deposited on a material, which irradiation properties are already well known (zr-based alloys serving as substrates). another objective is to test the behavior of the interface substrate/coating. structural changes at the interface caused by radiation might lead to issues related to adhesion of the coatings, cracking or excessive stress in the atf cladding. neutron irradiation will be always required to qualify materials for new reactor applications, however ion irradiation, as a low-cost and fast damage method, can serve as a screening method for evaluating new material concepts. with the use of ion irradiation, the development process of new materials for current and future reactor applications can be much faster and cheaper. 1relatively thin, limited by a penetration depth of the beams. 3. methodology and example results the proposed optimization method will be demonstrated on an example of irradiation of a thin ti2alc (max phase) coating deposited on zircaloy-4 alloy. zircaloy-4 is a standard binary zr-sn alloy used by the nuclear industry for decades. the max phase coating is proposed as a potential candidate material enhancing the safety features of the standard zr-based cladding. the ultimate goal is to irradiate the coating layer with multiple ion beams in such proportion that the sum of their respective damage profiles will result in an optimal damage distribution. the optimal damage distribution is defined as a flat distribution. that means a situation, where at each point of the depth of the target, there is the same irradiation-induced damage expressed by the quantity dpa per fluence [ dpa · cm2/ion ] . only the damage distribution in the coating layer of the target is taken into account, the shape of the damage profile in the substrate (zircaloy-4) is not considered. however, it should be noted that for the purposes of irradiation experiments, it is crucial for the beam to be able to penetrate the full depth of the coating, since the behavior of the interface between the coating and the bulk cladding is of a significant concern in the development of multi-component cladding materials. for that reason, the thickness of the coating needs to be optimized to fit the requirements of the test. in general, an arbitrary number of beams can be used for the optimization. there are two basic conditions the beams shall fulfill: first, at least one of the beams should be able to penetrate the full width of the coating layer. second, the bragg peaks of damage distributions created by individual beams should not overlap significantly. for demonstration purposes, it is assumed here that four different ion beams are to be combined, namely: fe (6 mev) and ni (6, 3.4, and 1.7 mev). each of these beams creates a unique damage profile2 when irradiating the target. these profiles are denoted as a(z), b(z), c(z), and d(z), respectively, and are functions of z the penetration depth inside the target. these profiles can be also represented as vectors a = {a1,a2, . . .anb}, b = {b1,b2, . . .bnb}, c = {c1,c2, . . .cnb}, and d = {d1,d2, . . .dnb}, where nb is the number of bins present in the coating part of the target. the aim is to find such parameters a, b, c, d fulfilling the fundamental condition a + b + c + d = 1, (1) so that the resulting damage profile fj = aaj + bbj + ccj + ddj, j ∈{1, . . .nb} (2) will be as flat as possible. 2damage profiles can be obtained using monte carlo simulation tools, e.g. srim [7] as in the case of this study. 44 vol. 14/2018 simulation of irradiation processes the flatness of the distribution can be described by its standard deviation. a small standard deviation corresponds to an even distribution of the defects, optimal (flat) distribution having the standard deviation equal to zero. the standard deviation of a damage distribution f is calculated as sdev = √√√√ 1 nb − 1 [ s2 nb − ( s1 nb )2] , (3) where nb is the number of bins present in the coating part of the target and s1 = nb∑ j=1 fj = nb∑ j=1 aaj + bbj + ccj + ddj, (4) s2 = nb∑ j=1 f2j = nb∑ j=1 (aaj + bbj + ccj + ddj ) 2 . (5) having defined the problem and the measure of quality of a solution of this problem, standard optimization tools can be employed. in this case, a proprietary implemented genetic optimization algorithm [8] was used to find an optimal set of parameters a, b, c, d. genetic algorithms work with five basic concepts: generation, population, individual, genotype, and gene. • each generation is represented by one population. • a population contains individuals, who represent possible solutions of the problem. • each individual is represented by its genotype a set of genes. • a gene can be a binary variable, an integer, a float or anything else. the algorithm iterates over generations with an aim to transfer the best individuals from the current generation to the following one. the quality of an individual is expressed by the value of its fitness function, the higher the fitness, the better the individual. a flowchart describing the genetic algorithm used for irradiation optimization is showed in figure 5. at the beginning, damage profiles are loaded from previously prepared files. (see number 1 in the figure 5.) an initial population is created using the method initialize_population as a set of random individuals. (nr. 2 in figure 5.) for each individual, the genotype (values a,b,c,d) is set as an array of random numbers from an interval 〈0, 1〉. eventually, the genotype is normalized so that the individual fulfills the condition (1). after creating the initial population, an iteration over the number of generations begins (nr. 3). following steps are executed in each iteration. • the quality of each individual in the population is calculated (nr. 4). individual’s quality is represented by the fitness function value. since the algorithm is searching for an individual with the highest fitness value, while our aim is to minimize the value sdev , the fitness of an individual is calculated as fitness = 1 sdev . (6) • a linear transformation is then applied on the fitness of each individual by subtracting the fitness of the worst individual in the population (nr. 5). the transformation improves the process of roulette selection, which will follow later. it ensures that high-quality individuals will be more likely to be selected for the process of crossing. • the index of the best individual in the population is saved. the best individual will be eventually artificially inserted into the new population (elitism). • a new population is created as an empty array (nr. 6). an iteration over a half of the number of individuals in each population (population_size/2) begins. following steps are executed in each iteration. • two individuals (parents) are randomly selected (nr. 7). • with the probability (1 crossing_rate) the parents are directly inserted into the new population (nr. 8), with the probability crossing_rate, they undergo the process of crossing (nr. 9). • in the crossing process, two new individuals (descendants) are created via random combination of genotypes of the parents. • after the crossing process, each gene of each descendant is, with a probability (1 mutation_rate) left as it is (nr. 10), or with a probability mutation_rate modified (nr. 11) by addition of a random value from the distribution g(x,y) = 0.01 √ −2 ln(x) sin(2πy), (7) where x and y are random values from an interval 〈0, 1〉. after adding a random value, it may happen that the value of a gene is negative. in that case, the negative value is replaced by zero. • the genotype of each descendant is normalized (nr. 12) so that he sum of all genes is equal to 1. • after the process of crossing (and mutation, in some cases), both descendants are inserted into the new population (nr. 13). • after population_size/2 iterations, the new population is completed (nr. 14). 45 václav šísl, martin ševeček acta polytechnica ctu proceedings 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 −0.03 −0.02 −0.01 0 0.01 0.02 0.03 mutation function g(x,y) = 0.01 √ −2 ln(x) sin(2πy) x y figure 4. two-dimensional mutation function used in the algorithm. • in the last step of the loop, the best individual from the previous population is artificially inserted into the new population and replaces the first individual created through natural selection. this method is called elitism (nr. 15) and is executed regardless of whether or not the best individual from the previous population is actually better than the one it replaces. the process of creating new populations based on crossing individuals from the previous one is repeated until a generation counter reaches a set value (nr. 16). after the loop terminates, the best individual is printed to the console (nr. 17) and the program is terminated. the described algorithm was applied on the previously mentioned case: irradiation of a thin layer of ti2alc coating deposited on zircaloy-4. this material is widely studied as a prospective candidate for the so-called accident tolerant fuel (atf) [9]. a visualization of the optimized damage profile obtained as a combination of four ion beams, fe (6 mev) and ni (6, 3.4, and 1.7 mev), is shown in figure 6. the ion beams (particles and their energy) were chosen based on available and widely used ion beams (e.g. [6]). the methodology and the algorithm are, however, very generally applicable. in fact, any combination of ion beams with any predefined energy can be used. the presented case serves only as an illustration of the presented method. the resulting parameters are: a = 0.05, b = 0.48, c = 0.22, and d = 0.25. the standard deviation was reduced by 98% compared to irradiation by a single beam (fe, 6 mev). the final values of parameters a,b,c,d can be then used to design an irradiation test of the particular fuel concept. the algorithm is very flexible and it allows users to optimize their own irradiation experiment. it can be easily modified based on available ions and their energies for a particular irradiation facility and particular material. there are certain simplifications compared to the real case. the recombination of defects and other temperature effects are neglected. furthermore, the beams act as independent meaning that the algorithm treats them separately without any interaction between them and neglects combined effects. it is, however, difficult to estimate the influence of these phenomena on the presented results. as for the issue of defect diffusion during irradiation, it would be convenient to perform the irradiation at decreased temperatures as it would make the process of diffusion slower. the diffusion related phenomena of radiation effects were studied by [10] who proposed a correlation that can be used to account for the different damage rates when using different types of beams. when using ions or protons imitating irradiation damage inside lwr reactors, the temperature of the target should be increased by 15–40 ◦c compared to the temperature in reactor. another approach might be to irradiate the sample with multiple beams at the same time not separately. in this case the defects are created uniformly, and the diffusion is slow due to a low defects concentration gradient in the material. this brings challenges to the operation of the accelerator and the experiment would be more complicated to perform. 4. discussion and conclusions it is believed that ion irradiation can serve as an inexpensive and fast verification method for application of new materials in the nuclear industry. the methodology of ion irradiation can be optimized to cope with the fundamental differences between the primary effects of ion and neutron irradiation as was shown in this paper. however, ion beam testing is unlikely to replace neutron irradiation in research reactors for the purposes of nuclear fuel licensing. the proposed method proved to deliver desired results a homogeneous distribution of irradiation damage in a layer, which is relatively thin, but thick enough to be used for testing atf multi-component cladding materials. on the other hand, a weakness of proposed solution is that only the issue of damage distribution was considered. all the other aspects, 46 vol. 14/2018 simulation of irradiation processes figure 5. flowchart describing the genetic algorithm used to optimize radiation damage distribution using multiple ion beams. 0 1 2 3 4 5 6 7 8 9 0 5 × 10−16 1 × 10−15 1.5 × 10−15 2 × 10−15 0 1 2 3 4 5 6 7 8 9 0 5 × 10−16 1 × 10−15 1.5 × 10−15 2 × 10−15 dp a pe r flu en ce [d pa ·c m 2 / io n] target depth [µm] ti2alc + zircaloy-4, dpa per fluence fe, 6.0 mev ni, 6.0 mev ni, 1.7 mev ni, 3.4 mev cladcoatdp a pe r flu en ce [d pa ·c m 2 / io n] target depth [µm] ti2alc + zircaloy-4, dpa per fluence optimized distribution cladcoat figure 6. damage profiles of four different ion beams irradiating 3 µm thick ti2alc coating layer deposited on zircaloy-4 slab (violet, green, blue and yellow) and their optimized combination (orange). such as damage morphology, temperature differences or defects diffusion have not been taken into account. it is, however, a perspective method enabling to imitate neutron irradiation using heavy ions. planned irradiation experiments will be optimized using the proposed method and the resulting damage distribution will be more representative for the real application inside nuclear reactor cores compared to current simple heavy ion irradiation. the optimization tool and its manual can be found on the github repository [8] and the application of this tool to real irradiation experiments with cr and crn coated cladding materials will be presented soon. the cr and crn coated zr-based alloys were irradiated also in the lvr-15 reactor in rez up to 1 dpa. one of the goals of the proposed research is to study the advanced cladding material irradiated at the research reactor and compare the results directly to the irradiation of the same cladding concept from an accelerator with the same irradiation damage. list of symbols atf accident tolerant fuel dpa displacements per atom pka primary knock-on atom sdev standard deviation 47 václav šísl, martin ševeček acta polytechnica ctu proceedings acknowledgements the support from student grant contest from czech technical university with code sgs16/252/ohk4/3t/14 is gratefully acknowledged. many thanks belongs to my supervisor ing. martin ševeček. references [1] g. s. was. challenges to the use of ion irradiation for emulating reactor irradiation. journal of materials research 30(09):1158–1182, 2015. doi:10.1557/jmr.2015.73. [2] g. was, r. averback. radiation damage using ion beams. comprehensive nuclear materials 1:195, 2012. [3] g. was, z. jiao, e. getto, et al. emulation of reactor irradiation damage using ion beams. scripta materialia 88:33 – 36, 2014. doi:http://dx.doi.org/10.1016/j.scriptamat.2014.06.003. [4] a. kohyama, y. katoh, k. jimbo. radiation damage study by advanced dual-ion irradiation methods. materials transactions 45(1):51–58, 2004. [5] y. serruys, m.-o. ruault, p. trocellier, et al. multiple ion beam irradiation and implantation: jannus project. nuclear instruments and methods in physics research section b: beam interactions with materials and atoms 240(1):124 – 127, 2005. accelerators in applied research and technology, doi:http://dx.doi.org/10.1016/j.nimb.2005.06.100. [6] s. taller, d. woodley, e. getto, et al. multiple ion beam irradiation for the study of radiation damage in materials. nuclear instruments and methods in physics research section b: beam interactions with materials and atoms 412:1–10, 2017. [7] r. stoller, m. toloczko, g. was, et al. on the use of srim for computing radiation damage exposure. nuclear instruments and methods in physics research section b: beam interactions with materials and atoms 310:75 – 80, 2013. doi:https://doi.org/10.1016/j.nimb.2013.05.008. [8] v. šísl. optimization tool for atf ion irradiation, github repository. https://github.com/vsisl/vyzkumny_ukol, 2017. [9] p. xu, e. lahoda, s. ray, et al. corrosion resistant coatings for zirconium-alloy cladding with improved accident tolerance. nuclear engineering and technology 48:16–25, 2016. [10] l. k. mansur. correlation of neutron and heavy-ion damage: ii. the predicted temperature shift if swelling with changes in radiation dose rate 78(1):156–160. doi:10.1016/0022-3115(78)90514-7. 48 http://dx.doi.org/10.1557/jmr.2015.73 http://dx.doi.org/http://dx.doi.org/10.1016/j.scriptamat.2014.06.003 http://dx.doi.org/http://dx.doi.org/10.1016/j.nimb.2005.06.100 http://dx.doi.org/https://doi.org/10.1016/j.nimb.2013.05.008 https://github.com/vsisl/vyzkumny_ukol http://dx.doi.org/10.1016/0022-3115(78)90514-7 acta polytechnica ctu proceedings 14:42–48, 2018 1 introduction 2 differences in effects of ion and neutron irradiation 3 methodology and example results 4 discussion and conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0046 acta polytechnica ctu proceedings 15:46–50, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app application of the modified magnetoelastic method and an analysis of the magnetic field tomáš kliera, tomáš míčkaa, michal polákb,∗, tomáš plachýb, milan hedbávnýc, roman jelínekc, filip bláhab,d a pontex, spol. s r. o., bezová 1658, 147 14 prague 4, czech republic b department of mechanics, faculty of civil engineering, czech technical university in prague, thakurova 7, 166 29 prague 6, czech republic c freyssinet cz a. s., zápy 267, 250 01 brandýs nad labem, czech republic d department of concrete and masonry structures, faculty of civil engineering, czech technical university in prague, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: polak@fsv.cvut.cz abstract. in the technical praxis there is very often a need of axial force determination in the important structural elements of a building during its construction or operational state. the paper presents practical application of the new approach based on the magnetoelastic principle, especially aimed on experimental evaluation of the axial forces in the prestressed reinforcement on prestressed concrete structures. described approach is usable not only for newly built structures but in particular for existing ones. the experiment was realized on a prestressed concrete beam dismantled from a bridge that was put out of operation. the influence of a simulated defect of the reinforcement on its magnetoelastic properties has also been investigated. during result evaluation, the knowledge based on the theoretical analysis of the electromagnetic field was used. keywords: concrete, tensile force, magnetoelastic method, prestressed strand, prestressed cable, prestressed wire, elastomagnetic sensor. 1. introduction in civil engineering practice, five experimental techniques are usually used for evaluation of axial tensile forces, the direct measurement of the force by a preinstalled load cell, the approach based on a strain measurement with strain gauges, the vibration frequency method [1, 2], the force determination in a flexible structural element based on the relation between the transverse force and the caused transverse displacement [3, 4], and the magnetoelastic method [5– 10]. all mentioned methods have their advantages and disadvantages that are discussed in more detail in the reference [7] and [8]. only three of them are usable for new experiment on an existing structure that is in service. the first one is the vibration frequency method that is applicable only to structural elements with relatively long free vibrating length. the uncertainty of the specified force is moreover significantly influenced by the element bending stiffness and by the element boundary conditions especially if they are complicated and vague [1, 2]. the second one is method based on the relation between the transverse force and the transverse element displacement that is usable only for significantly flexible elements with a relatively long free length. due to the transverse displacement the tensile force in the considerably short element during the experiment substantially increases even in the range of tens of percent [3, 4]. the third one is new approach based on the magnetoelastic method [5–7]. according to our opinion, the new approach is the most suitable procedure for tensile stress evaluation in prestressed elements on existing structures that are in service. the basic aim of the project is to research and develop a new approach based on the “magnetoelastic” physical principle. the approach is designed for the measurement of axial forces in structural elements made of ferromagnetic materials (in steel prestressed rods, wires and cables) on existing or newly constructed building structures. the research solved by our team seeks to focus primarily on solving the task of “determining the actual size of the force in a prestressed concrete element on an existing structure” which is very common in technical practice but is often not resolvable by any other method used nowadays. this paper presents one of the typical problems that have been solved so far [8–10]. in the conclusion, the properties and risks of this approach are summarized. 2. theory – physical principle elastomagnetic resp. magnetoelastic principle was explained in more detail in the sources [5, 6]. here are briefly summarized the basic facts: • the mechanical state and magnetic properties of the ferromagnetic material interact. 46 http://dx.doi.org/10.14311/app.2018.15.0046 http://ojs.cvut.cz/ojs/index.php/app vol. 15/2018 application of the modified magnetoelastic method... • the magnetic properties of the particular material define the relationship between the magnetic field shape described by the magnetic field intensity “h” [a.m−1] and the shape of the magnetic induction field “b” [t] (also named the magnetic flux density). for non-ferromagnetic materials the dependence is elementary, while the dependence for ferromagnetic ones is very complex, influenced by a number of parameters (temperature, mechanical stress, history of magnetic and mechanical stress for example). • the magnetic field intensity “h” can be monitored locally, for example, by hall sensors. the field of magnetic induction “b” may be monitored by means of a coil that integrates the normal component of the magnetic induction field over the area of individual coil turns. 3. theory – physical principle figure 1 shows a diagram of a fully equipped elastomagnetic sensor (hereinafter em sensor). fundamental sensor components are a controlled magnetic field source (the primary coil in this case), a sensor of magnetic field intensity “h” in a measured crosssection (the system of hall’s sensors, the secondary coil 2), a sensor of magnetic flux that is closely related to the magnetic induction “b” in the measured section (the secondary coil 1) and em sensor protection against magnetic influences from its surroundings (steel shield). figure 1. diagram of a fully equipped elastomagnetic sensor (em sensor). the function and principle of hall´s sensors were explained in more detail in the reference [6]. 3.1. maximal configuration the reader of this paper certainly assumes that a fully featured sensor offers the greatest possibilities to increase accuracy and reduce uncertainties in determining the axial force in the measured prestressed element. but the reader will also not be surprised that the fully equipped sensor is spatially larger, that may limit its applicability, and also more complicated. this complexity affects its higher production time, the higher usage requirements for this sensor in civil engineering practise and as well as the higher demands on the equipment of the used measuring system. 3.2. minimal configuration here is described an opposite extreme of the em sensor – the minimally featured one. the minimalist em sensor consists of a primary coil and a secondary coil 1 only. in this case, the intensity of the magnetic field “h” is determined indirectly from a completely different physical quantity, from the current flows through the primary coil. the greatest risk of using such sensor results from this approach. any change in the magnetic surroundings around the sensor (a removal of a massive steel falsework after concrete hardening for example) causes completely “silently” the substantial or fatal changes in sensor parameters. the sensor’s steel shielding reduces the range of this effect, the level of this reduction depends on the quality of the em sensor shielding, nevertheless it is never one hundred percent. for example, the minimum sensor configuration is completely unusable when applied to steel-fibre concrete. the em sensors and their appropriate equipment that are standardly used at present [8–10] evaluate measured prestressed forces in a relatively simple way. however, these standard em sensors, that have hitherto been used in civil engineering practice, are composed from the primary and secondary coils only and that is the minimal possible configuration of the em sensor as it is mentioned in this chapter. 3.3. measurement of magnetic field intensity if we are not satisfied with the indirect measurement of the magnetic field intensity “h” described in the previous paragraph, it is possible to use the hall sensors for this purpose. well-calibrated hall sensors allow very accurate determination of “h” in close proximity to the studied object, for example even in the gaps between the strands or between the prestressed wires in the experimentally investigated prestressed cable, and they are particularly advantageous when the observed magnetic field is asymmetrical. one disadvantage of this em sensor configuration is the fact that modern hall sensor are integrated semiconductor circuits that are not, in principle, reliable for long-term use in exposed surroundings. an alternative to hall sensors is to add a secondary coil 2 to the sensor construction. this coil, as well as the concentrically located secondary coil 1, observes a change in the magnetic flow passing through its turns. the mutual difference between magnetic flows measured by the coil 2 and the coil 1 indicates a change in flow passing through the annulus between turns of both secondary coils. in this space there are only non-ferromagnetic materials and therefore the magnetic flux can be readily recalculated to the average normal component of the magnetic field intensity “h” 47 t. klier, t. míčka, m. polák et al. acta polytechnica ctu proceedings in this annulus. the disadvantage of this sensor arrangement is that the intensity of the magnetic field is determined in comparison with the hall sensors at a greater distance from the studied prestressed elements and also that the secondary coil 2 increases the dimension of the sensor transverse cross-section. on the other hand, the advantage of this em sensor configuration is that the output of the em sensor is integrated from the larger area and therefore it is more resistant to the effects of local anomalies in the magnetic field and the application of secondary coil 2 also allows to not use the hall probes within the em sensor and it positively affects the reliability of the em sensor in its long-term use. 3.4. optimal configuration of the em sensor for existing structures the so far gained experience by em sensor applications on existing structures shows that it seems optimal to assemble an em sensor from all the parts shown in figure 1. its complete configuration allows the best available calibration of an em sensor under specific conditions. if the long-term use of the retrofitted em sensors is assumed, then it is possible to dismantle the hall probes from the em sensor after its initial calibration, and then process only the outputs of a pair of secondary coils. otherwise, for a one-off experiment, it is possible to omit the secondary coil 2 installation because there is no use for it. 4. experiment on a real structure the experiment was carried out on the prefabricated prestressed concrete beam of the type ka-67, which was dismantled from a bridge put out of service and which did not serve the original purpose for years. thanks to this fact, the experiment could be done without minimizing damage of the concrete part of the beam. it was also possible to carry out test in which an influence of a simulated defect of the studied cable on its magnetoelastic properties has been investigated. the force in the prestressed steel cable was intentionally released by means of the gradual cut off prestressed wires forming cable. this experiment was focused on the following problematic areas: verification of real spatial requirements for an em sensor additional installation on a prestressed cable, optimization of the geometric arrangement of the em sensor to real conditions and theoretical modelling of the electromagnetic field of the sensor taking into account the general geometric arrangement of the measured cross-section of the studied prestressed reinforcement. the theoretical model created in the software ansys maxwell 3d that corresponds to the experiment arrangement and conditions is shown in figure 5. the design of the applied em sensor has been optimized for the particular conditions of one-off experiment realized on the “small” in-situ cable, see figure 2. the production of the secondary coil in-situ. figure 3. the holder for the set of four hall sensors with the embedded secondary coil 1 positioned on the studied cable. figure 2. a cable composed of only a few prestressed wires allowed the use of a portable primary coil to emit a sufficiently intense electromagnetic field, see figure 2. only the secondary coil 1 was produced in-situ, see figure 1 and figure 3. the set of hall sensors located on the holder stabilizing the ideal mutual position of all sensor parts relative to each other was also demountable and portable. the steel shield with regard to detailed magnetic field monitoring in the measured cable cross-section by four hall probes was not necessary. the condition of successful application of the elastomagnetic measuring method is the detailed knowledge of the composition of the electromagnetic field generated by the primary coil in the measured cross-section of the studied prestressed element. based on the analysis of the results obtained from the theoretical model (see figure 3), it is possible to create a functional relationship between the real measured variables and the variables necessary for the evaluation of the solved prestressing force or the stress in the studied prestressed reinforcement. for example, we need to know the average value of the magnetic field intensity in the 48 vol. 15/2018 application of the modified magnetoelastic method... figure 4. the complete em sensor on the cable. figure 5. the theoretical model of the em sensor positioned on the studied prestressed cable – only half of the model visualization. cross-section of the studied prestressed element, but we can only measure it locally at the positions of hall sensors. in figure 7 there is a graphical representation of the results obtained during the realized experiment. the red lines with the “+” node symbols show the measured values of the average stress in the cable. they differ by the evaluation curve used for determination of the results. davg is the mean curve obtained from available results evaluated by previously realized experiments carried out for three different types of prestressed wires [6, 7] and d1 is the most appropriate evaluation curve obtained for the prestressed wire most similar to that used in the tested structural element. the black and blue curves show the computed values of the average stress in the studied cable according to different computational models of the prestressed figure 6. the theoretical model of the em sensor with the studied prestressed cable – magnetic field intensity in the cross-section of the secondary coil 1. figure 7. 7 development of the axial tensile force in the cable during the gradual prestressed wires cut off. concrete beam. the black curve corresponds to a model where zero deformation of concrete is assumed during the gradual cutting off of the wires. the blue thin curve corresponds to the assumption of uniform elastic deformation in concrete. in addition, the blue thick curve includes a local change in the field of deformation due to the demolition work performed to allow the installation of the em sensor. 49 t. klier, t. míčka, m. polák et al. acta polytechnica ctu proceedings the very good match between the average stresses measured and evaluated according to the curve d1 (thick red curve in figure 7) and the non-linear static model of the observed structural element (thick blue curve in figure 7) is obvious. the more significant differences in these two curves occurred only in the area where almost all cable wires are cut off. the deviation between these two curves is probably caused by different stresses than axial. in the case of the described experiment, it is probably due to the weight of the primary coil acting in the transverse direction on the remaining uncut wires. 5. conclusions the presented experiment confirmed the applicability of the developed new measuring approach based on the “magnetoelastic” physical principle to real existing structures. accuracy and uncertainty of results obtained by this approach are significantly affected, as other methods, by the quality of the experiment realization, but also by the quality of the used evaluation curve. obviously, the highest quality of the evaluation curve can be obtained from a prestressed reinforcement sample taken directly from the investigated structure, but in most real cases it is difficult or totally impossible. therefore, it is important to gradually create the widest possible database of evaluation curves for prestressed elements commonly used in the building industry, both today and in the past. acknowledgements the results presented in this paper are outputs of the research project fv 30457 “utilization of a magnetoelastic method for increasing the reliability and durability of existing and newly built prestressed concrete structures” supported by ministry of industry and trade of the czech republic. references [1] m. polák, t. plachý. determination of forces in roof cables at administrative center amazon court. procedia engineering 48:578–582, 2012. [2] m. polák, t. plachý. experimental evaluation of tensile forces in short steel rods. applied mechanics and materials 732:333–336, 2015. [3] p. fajman, m. polák. measurement of structural cable of membranes. in proceedings of the 50th annual conference on experimental stress analysis, vol. 50, pp. 61–64. 2012. [4] p. fajman, m. polák, j. máca, t. plachý. the experimental observation of the prestress forces in the structural elements of a tension fabric structure. applied mechanics and materials 486:189–194, 2014. [5] t. klier, t. míčka, m. polák, et al. modern method for measurement of the axial forces in elements of building structures. in proceedings of the conference betonářské dny 2016, vol. 23. čbs, 2016. [6] t. klier, t. míčka, t. plachý, et al. the verification of a new approach to the experimental estimation of tensile forces in prestressed structural elements by method based on the magnetoelastic principle. matec web of conferences 107(00015):1–6, 2017. [7] t. klier, t. míčka, m. polák, et al. the in situ application of a new approach to experimental estimation of tensile forces in prestressed structural elements by method based on the magnetoelastic principle. in proceedings of the 55th conference on experimental stress analysis 2017, vol. 55, pp. 122–132. technical university of košice, 2017. [8] a. m. sarmento, a. lage, e. caetano, j. figueiras. stress measurement and material defect detection in steel strands by magneto elastic effect. comparison with other non-destructive measurement techniques. in proceedings of the 6th international conference on bridge maintenance, safety and management iabmas 2012, pp. 914–921. iabmas, 2012. [9] m. chandoga, p. fabo, a. jaroševič. measurement of forces in the cable stays of the apollo bridge. in proceedings of the 2nd fib congress, pp. 674–675. fib, 2006. [10] h.-j. wichmann, a. holst, h. budelmann. magnetoelastic stress measurement and material defect detection in prestressed tendons using coil sensors. in proceedings of 7th international symposium on nondestructive testing in civil engineering ndtceš09, p. 6. laboratoire central des ponts et chausses, 2009. 50 acta polytechnica ctu proceedings 15:46–50, 2018 1 introduction 2 theory – physical principle 3 theory – physical principle 3.1 maximal configuration 3.2 minimal configuration 3.3 measurement of magnetic field intensity 3.4 optimal configuration of the em sensor for existing structures 4 experiment on a real structure 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0022 acta polytechnica ctu proceedings 7:22–28, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app validation and implementation of failure parameters in integrated simulations for short fibre reinforced polypropylene anna kalteisa, ∗, martin reitera, michael jerabekb, zoltán majora a institute of polymer product engineering, johannes kepler universität linz, altenberger straße 69, linz, austria, 4040 b borealis polyolefine gmbh, st.-peter-straße 25, 4021 linz ∗ corresponding author: anna.kalteis@jku.at abstract. nowadays short fibre reinforced polymers are often used in load carrying structural parts. compared to continuous fibre reinforced polymers they exhibit a more complex morphology. hence the determination of the strength is a difficult but important task. therefore this was the objective of this research. the strength of short fibre reinforced polymers was numerically determined for low-speed to high-speed strain rates for specimens with different fibre orientations. for the failure modelling the micromechanical approach “first pseudo grain failure” in digimat was used. the parameters for the material and failure description were determined with the reverse engineering method. integrated finite element simulations were performed to validate the material and failure models by tensile and bending tests with different specimens. the comparison of the results of the experiments and simulations showed low deviation. keywords: integrated simulation, short fibre reinforced polymer, failure prediction, first pseudo grain failure, reverse engineering. 1. introduction the prediction of the stiffness of a fibre reinforced material is not a difficult task anymore because there is appropriate software available. one is the software digimat mf. there only few input parameters are needed to calculate the stiffness [1]. however the prediction of the strength of a material especially for short fibre reinforced polymers is not trivial. to perform simulations is very common nowadays, therefore in this research the strength of short glass fibre reinforced polypropylene is determined by integrated finite element simulations. there are many different approaches how to predict the strength of a discontinuous fibre reinforced polymer. for example there is an analytical model where the strength of fibre reinforced composites can be calculated without any numerical software [2]. furthermore there are several continuum mechanics based models [3], [4], [5], [6], [7], [8], [9] and micromechanics based models [10], [11], [12], [13]. in the paper also a micromechanics based model is used namely the “first pseudo grain failure”(fpgf) model [1]. a short description is given in section 4 of this paper. the investigated material is a short glass fibre reinforced polypropylene. the glass fibre content is 32 w%. because of the short fibre reinforcement the fibre distribution and orientation in the different specimens have to be investigated. the fibre orientations were determined by the computer tomographic method. the raw data from the computer tomographic measurements of the fibre orientation were provided by fh-oö (wels, austria). therefore a description of the ct-measurements is not given here but in [14]. the experimental data were provided by borealis polyolefine gmbh (linz, austria). standardized tensile (iso 527) and bending tests (iso 178) were performed with the corresponding strain rate and specimen. specimens with different complexity of fibre distribution and orientation were used. specimens with a uniform fibre distribution and orientation (ud multi tool) are used for the reverse engineering. for the validation specimens with a more complex fibre distribution and orientation are used. a description of the used specimens is given in section 2. the modelling of the material behaviour is an other part of this paper. elastic-plastic (low-speed strain rates) and elastic-viscoplastic (low-speed and highspeed strain rates) material models are generated. the needed material parameters are reverse engineered based on experimental data of tensile tests. in section 4 the determination of an appropriate failure criterion is described. because the polymer is reinforced with short fibres classical failure criteria (e.g. tsai-hill) cannot be simply used because they are just valid for continuous fibre reinforced composites [1]. hence as already mentioned above the “first pseudo grain failure” (fpgf) model has to be used. the failure parameters are determined with the method of reverse engineering with data of tensile tests. the results of the material and failure modelling are validated with tensile tests with specimens with different fibre distribution and orientation as the reverse engineered ones. further validation follows by per22 http://dx.doi.org/10.14311/app.2017.7.0022 http://ojs.cvut.cz/ojs/index.php/app vol. 7/2017 forming bending tests. the appearing difficulties in the simulation of the bending test have to be solved and are described in section 5.3. the validation of the material and failure model is done for low-speed and high-speed strain rates. in this paper only the results for the low-speed strain rates are discussed. 2. specimens as already mentioned different types of specimens are used for the experiments and simulations. an overview of the used specimens is given in fig. 1. figure 1. overview of different specimens. different specimens are used because of their different complexity of fibre orientation. the ud-multitool was designed by borealis polyolefine gmbh (linz, austria). the specimens, which can be milled out of the tool have a highly uniform fibre distribution. the specimens 0◦, 45◦ and 90◦ are relevant for this paper. because of the uniform fibre distribution these specimens are used for the reverse engineering of the material and failure parameters. for the validation of the material and failure models the iso-mps specimen and the plate specimens are used. the plate specimens were milled out of a plate at different positions and under the angles 0◦, 45◦ and 90◦. these specimens do not show a uniform fibre orientation through the thickness as the ud-multitool specimens. 3. material modelling for the material modelling the software digimat mf 6.0.1 is used. for the glass fibre a linear elastic, isotropic material model is chosen. therefore only the elastic modulus and the poisson’s ratio have to be defined. for the elastic modulus and poisson’s ratio the following values are used: e = 72000 mpa, ν = 0.2 the aspect ratio (ar) of the fibres is set to 35 [15]. an elastic-plastic (ep) and also an elastic-viscoplastic (evp) material model are generated for the matrix. the ep and evp material models for the matrix need more input parameters than an elastic one. the j2plasticity model is chosen (eq. 1). for the hardening the exponential and linear law is used (eq. 2). where r(p) is the hardening stress, p is the accumulated plastic strain, k is the linear hardening modulus, r∞ is the hardening modulus, m is the hardening exponent [1]. σeq = √ j2(σ) (1) r(p) = kp + r∞[1 − exp(−mp)] (2) for the elastic-viscoplastic model some additional parameters for considering the viscoplasticity have to be determined. a law for the accumulated plastic strain rate ṗ has to be chosen. in this paper the current yield norton law is chosen (eq. 3). there the ṗ is dependent on the initial yield stress σy, the viscoplastic stress f, the hardening stress r(p), the viscoplastic coefficient η and the viscoplastic exponent m. ṗ = σy η ( f σy + r(p) )m (3) a detailed description is given in [1]. the above described parameters are reverse engineered with the software digimat mx 6.0.1. with this procedure the parameters for an artificial matrix are generated. they are artificial because the real behaviour of the matrix (e.g. micro-cracking) and the influence of the fibre on the matrix (e.g. fibre-matrix debonding) are inherently considered. these influences cannot be modelled by using the parameters for the neat matrix material [16]. in the reverse engineering procedure experimental stress-strain curves are needed where the microstructure (fibre orientation) is known. the parameters of the matrix are fitted to the experimental curves in an iterative optimization process [1]. the results of the material modelling for the ep and evp material models for different strain rates are summarized in table 1 and table 2. 4. failure modelling the intralaminar failure criteria, which are available in digimat are only valid for a lamina layer with unidirectional aligned fibres [1]. they can be non-interactive, interactive or failure mode based [17]. the failure criterion, which is used is tsai-hill 3d transversely isotropic (stress based, interactive). this failure criterion assumes that the strength in tension and compression is identical and the material is isotropic in the 2-3 plane [1]. direction 2 and 3 are the directions transverse to the fibre orientation (direction 1). for the description of the failure criterion three input parameters have to be defined: 23 a. kalteis, m. reiter, m. jerabek, z. major acta polytechnica ctu proceedings strain rate [s−1] 10−4 10−3 10−2 young’s modulus [mpa] 1800 1800 1800 poisson’s ratio [-] 0.42 0.42 0.42 yield stress [mpa] 4 4 4 hardening modulus [mpa] 10.3 13.3 14.2 hardening exponent [-] 154.5 208.2 188.3 linear hardening modulus [mpa] 22.6 19.3 32.2 table 1. reverse engineered elastic-plastic material parameters for low-speed strain rates. strain rate [s−1] low-speed (10−4, 10−3, 10−2) high-speed (0.1, 1, 14) young’s modulus [mpa] 1800 1800 poisson’s ratio [-] 0.42 0.42 yield stress [mpa] 3 4 hardening modulus [mpa] 12 9 hardening exponent [-] 145 85 linear hardening modulus [mpa] 22 60 viscoplastic coefficient [mpa*s] 51 71 viscoplastic exponent [-] 4 6 plastic strain multiplier [-] 1.5 1.9 table 2. reverse engineered elastic-viscoplastic material parameters for low-speed and high-speed strain rates. • axial tensile strength x (tensile strength in fibre direction) • in-plane tensile strength y (transverse tensile strength) • transverse shear strength s the failure criterion is defined as follows [1]: fa(σ) = σ211 x2 − σ11(σ22 + σ33) x2 + σ222 + σ233 y 2 +( 1 x2 − 2 y 2 ) σ22σ33 + σ212 + σ213 s2 +( 4 y 2 − 1 x2 ) σ223 (4) because a short fibre reinforced polymer with nonperfectly oriented fibres is used and not a lamina layer with perfectly aligned fibres the “first pseudo-grain failure” (fpgf) method has to be used. with this method the failure criteria for laminates can also be used for composites with complex fibre orientation distributions because the short fibres are divided into regions with uniform fibre orientations [1]. this is illustrated in fig. 2. the strength values are reverse engineered based on the ep material model. also here the data of the tensile tests of specimens of the ud multi tool are used. in fig. 3 the strength parameters of the failure figure 2. illustration of the first pseudo-grain failure method [16]. modelling for different strain rates are illustrated. it can be seen that the higher the strain rate the higher the strength is. 5. integrative simulation 5.1. interface digimat cae for the validation of the material model integrative digimat-abaqus fem simulations are performed. for these simulations an interface is needed. thus the software digimat cae 6.0.1 is used. with this software the material properties are calculated depending on the fibre orientation either before or during the simulation. this depends on which so24 vol. 7/2017 figure 3. illustration of strength parameters for different strain rates. lution is chosen in digimat cae. in the following the solutions are described [1]: • micro solution: the material properties are calculated at every iteration and not kept constant. as a result the simulation takes very long. • hybrid solution: the material properties are precomputed before the simulation. this enables that the simulation is sped up. with this method so called hybrid parameters are computed during the generation of the material input file. integrative simulations (tensile test) with the two different solution methods are performed. the results show that there is nearly no difference between micro and hybrid solution (fig. 4). the hybrid solution is chosen for the further simulations because of the lower computation time. figure 4. comparison of hybrid and micro solution method. 5.2. simulation model of tensile test as the specimens show a fibre orientation distribution through the thickness this distribution has to be discretised for the simulation model. the discretised distribution can be modelled with a layer build-up of the specimen (fig. 5). the layers can be modelled by partitioning the specimen through the thickness and assigning a material model with a specific fibre orientation to every partition. the number of layers can vary. the higher the number of layers the more accurate the fibre distribution is described. furthermore the higher the number of layers is the higher the computation time because more different material models have to be assigned. for the tensile tests a 4 layer model is used. therefore for the simulation of a tensile test a 4 layer model with the hybrid solution method is used. in this simulation model no problems occur. figure 5. simulation model with 4 layers. 5.3. simulation model of bending test because there are no problems in the simulation model of the tensile test these settings are also applied for the bending tests. therefore a bending test with a 4 layer model and with the hybrid solution method is modelled. in contrast to the tensile test simulation model, the bending simulation model shows various complications by using exactly the same simulation approach. therefore the settings of the simulation model of a tensile test cannot be directly transferred to the simulation model of the bending test. the simulation model of the bending test has to be modified. the problems in the simulation model of the bending test (hybrid solution, 4 layers) and the belonging solutions are explained in table 3. figure 6. specimen with global fibre orientation. the soft contact, which is used is a normal behaviour with exponential law between the master and slave surface. this contact is implemented in abaqus. two parameters (clearance and pressure) have to be defined. with this contact the pressure of the stamp 25 a. kalteis, m. reiter, m. jerabek, z. major acta polytechnica ctu proceedings problem solution 1 fibre orientation does not rotate with specimen applying global fibre orientation (fig. 6) 2 elements break under pressure soft contact between specimen and stamp (fig. 7) 3 “kink” in curve (fig. 8) 10 layer model instead of 4 layer model 4 considering of different behaviour under compression [18] adjustment of hardening modulus for the upper 5 layers (fig. 9) 5 no valid hybrid parameters could be deter-mined for some layers using micro solution 6 specimen breaks at a too low stress adjustment of failure parameters table 3. description of the problem and the corresponding solution for the bending simulation model. figure 7. illustration of position of applied soft contact. figure 8. illustration of kink in curve. figure 9. adjustment of material behaviour under compression; a) illustration of layers for the adjustment; b) red curve: stress-strain curve without adjustment, green curve: stress-strain curve with adjustment. is distributed over a larger area of the specimen [19]. therefore the stress concentrations on the specimen are lowered and as a result the specimen does not break under pressure anymore. the parameters are investigated by performing simulations of a bending test to find the most appropriate ones to solve the problem that the elements break under pressure. the reason for using a 10 layer model for the bending test is that the stress differences in the 10 layer model are lower than in a 4 layer model. this is because in the 10 layer model there is a smoother transition between the different fibre orientations. therefore the “kink” in the curve vanishes. unlike in the simulation model of the tensile test in the simulation model of the bending test the hybrid solution cannot be used. the reason for this is that in some layers no valid hybrid parameters could be determined by the software. therefore the micro solution has to be chosen. the failure parameters are reverse engineered for the ep material model. but for the bending tests the evp material model is used in order to consider the strain rate dependency. therefore the reverse engineered failure parameters may not fit as well for the evp material model as for the ep material model. as a result the parameters are fitted as it is described in section 6.2. 6. results 6.1. results of tensile test fig. 10 shows the result of the tensile test for a lowspeed strain rate. the material model is the elasticplastic one with strain rate of 10−3s−1. as the strain rate for the simulation is implemented in the elasticplastic material model also for the low-speed strain rates the dynamic explicit solver can be used to speed up the simulation. also an uncoupled simulation was performed in digimat mf. a difference in the failure stress between the simulations performed in abaqus 6.14 (integrative simulation) and digimat mf can be seen. 26 vol. 7/2017 figure 10. comparison of stress-strain curves of experiment, simulation performed in digimat mf 6.0.1 and simulation performed in abaqus 6.14. the reason for this is that the specimen shows a stress concentration in the region of the shoulder (fig. 11). figure 11. illustration of location of stress concentrations; a) specimen before failure, b) failed specimen. the notch effect at the specimen typically does not appear in reality. in reality the iso-mps specimen should break in the middle. in the simulation this cannot be realized in any specimen. because the simulation model itself cannot be modified to consider the notch effect, the failure parameters have to be fitted. therefore the parameters are fitted in this way that the notch effect can be considered. this fitting is performed by lowering the strain rate dependency by a power of 10. this means that the failure parameters for the strain rate of 10−2 s−1 are used for the elasticplastic material model for the strain rate of 10−3 s−1. in the following figure the results of the experiment and the simulation with the fitted failure parameters are shown. figure 12. comparison of stress-strain curves of experiment and simulation performed in abaqus 6.14 with fitted failure parameters. 6.2. results of bending test the bending test simulations are performed with the elastic-viscoplastic low-speed material model. because of the viscoplastic behaviour the strain rate has to be defined in abaqus. the static solver is used. in the following the result of the bending test for a specimen with misaligned fibres with the not fitted failure parameters is shown (fig. 13). as it can be seen the specimen breaks at a lower force than the experiment. one reason for this is described in section 5.3. figure 13. comparison of force/thicknessdisplacement curves of experiment and simulation. therefore the in-plane tensile and the transverse shear failure parameters are fitted. the axial tensile strength is kept constant because it has the most influence on layers with fibres, which are nearly perfectly aligned. in table 4 the previous and the fitted parameters are listed for the two different parameters. the difference between the previous and the fitted value is not very big. nevertheless these little increased values have a huge influence on the global result, which is illustrated in fig. 14. strength parameter previous value fitted value in-plane tensile strength 22.83 mpa 25.32 mpa transverse shear strength 13.38 mpa 13.60 mpa table 4. previous and fitted strength parameters. figure 14. comparison of force/thicknessdisplacement curves of experiment and simulation with fitted strength parameters. 27 a. kalteis, m. reiter, m. jerabek, z. major acta polytechnica ctu proceedings 7. conclusion with the knowledge of the specimen’s fibre orientation simulations can be set up, where this fibre orientation is considered. material and failure modelling was done for specimens with different fibre orientations and for low-speed to high-speed strain rates. it was shown that it is possible to determine failure in tensile and bending tests with low deviation (fig 15). figure 15. illustration of the deviation between failure force in the experiment and in the simulation. many influences have to be considered in a bending test (e.g. applying a global fibre orientation). little different values of failure strength parameters have a huge influence on the failure strength of the specimen. acknowledgements i would like to thank dr. michael jerabek for the possibility to do this research in cooperation with borealis polyolefine gmbh. references [1] digimat documentation 6.0.1. 2015. 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[12] a. krairi, i. doghri, g. robert. multiscale high cycle fatigue models for neat and short fiber reinforced thermoplastic polymers. international journal of fatigue 2016. [13] w. lutz, m. dong, k. zhu, s. schmauder. modeling of damage in fibre and particle reinforced composites, in damage and its evolution in fiber-composite materials: simulation and non-destructive testing. isd-verlag, 2006. [14] j. kastner, e. schlotthauer, d. angermaier, g. zitzenbacher. quantitative messung von faserlängen und -verteilung in faserverstärkten kunststoffteilen mittels µ-röntgen-computertomographie. dgzfp-jahrestagung 2007. [15] r. teschner. glasfasern. springer vieweg, 2013. [16] m. reiter. development, implementation and validation of an integrated simulation methodology for discontinuous fiber reinforced composites. ph.d. thesis, institute of polymer product engineering, johannes kepler university, linz. 2014. [17] i. m. daniel, o. ishai. engineering mechanics of composite materials. oxford university press 2006. [18] a. hartl. mechanical behavior of short glass fiber reinforced and particle modified polypropylene, influence of material anisotropy and loading mode. ph.d. thesis, institute of polymeric materials testing, johannes kepler university, linz. 2014. [19] abaqus analysis user’s guide, 6.14. 2014. 28 acta polytechnica ctu proceedings 7:22–28, 2017 1 introduction 2 specimens 3 material modelling 4 failure modelling 5 integrative simulation 5.1 interface digimat cae 5.2 simulation model of tensile test 5.3 simulation model of bending test 6 results 6.1 results of tensile test 6.2 results of bending test 7 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0043 acta polytechnica ctu proceedings 11:43–48, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app netgraph – an efficient tool for periodic timetabling and capacity planning zdeněk michl∗, michal drábek, rudolf vávra department of logistics and management of transport, faculty of transportation sciences, ctu in prague, czech republic ∗ corresponding author: michlzde@fd.cvut.cz abstract. a netgraph is a graphical representation of a periodic timetable concept for given spatial or modal scope. it adds a third dimension to the network plan – time. despite no strictly prescribed rules, a common practice notation has been settled over time. the authors propose an extension of a netgraph for planning of periodic railway capacity and operation, including related notation rules. keywords: edge, netgraph, node, notation, periodic timetable, periodic capacity. 1. introduction a netgraph, also called interval graphic, is a graphical representation of a periodic (interval) timetable concept of public transport lines in a designated geographical or political territory, for a transport segment or a production unit. it displays at first traffic and transport coherence of the individual lines in transfer nodes and their service function on the territory (number of stops). netgraph was made up of a combination of a network plan, as usually used in urban and local public transport, and a timetable, as used in railway timetabling. while the former reproduces the course of the lines and partially the intervals in the entire network, the latter shows in detail the scheduled traffic on a part of the network (mostly a single railway line). the netgraph adds a third dimension to the two-dimensional network plan – time schedule. thus, in a netgraph, periodic timetable with small to medium intervals in an area can be clearly illustrated. there are no strictly prescribed rules for creation of netgraphs, but over time a common practice notation has been settled. since the level of detail, the purpose of creation and highlighted properties in netgraphs varies, it is always necessary to provide a legend of line (edge) types and colours and time notation used in it. a netgraph may be created as an output of a specialised transport planning software package as german fbs – fahrplanbearbeitungssystem (fbs/iplan) [1], swiss viriato [2] and others, enabling an overview of the constructed lines and connections at a glance for comparison. a basic netgraph can even be constructed using a spreadsheet office suite or another graphic tool. a netgraph consists of following elements: • nodes, including their names, • edges, including the timetable notation for each public transport line and number of stops, • technological processes, • a legend. further, the authors propose an extension of a netgraph for planning of periodic railway capacity and operation. for such purpose, it is necessary to establish rules for notation of passing through defined station at a certain time, for alternative routing of a train path, for a train path for one direction only, and for crossing of the opposite trains on a single-track line. a proposal of related rules is presented and discussed consequently. 2. common practice in netgraph notation there emerged two main producers of netgraphs – irfp. e. k. (institut für regional-und fernverkehrsplanung), and sma und partner ag. both companies provide timetabling software tools. their rules for netgraph notation are mostly equal, except some differences (e.g. as in figure 9 by sma and as in figure 10 by irfp). both companies publish on their websites a short introduction into purpose and notation of a netgraph. irfp uses a term "interval graphic", sma uses "netgraph". the authors of this paper have decided for the term "netgraph" for the sake of brevity. irfp notes that a netgraph displays at least • geographical offer: (public transport) "line network and line management, main railway stations, operated access points per line • temporal offer: interval frequence per line as well as offer density on certain lines through overlap, arrival and departure times of the clocked lines in main railway and intermediate stations, determining of lines and transition periods in the stations • qualitative offer: long-distance traffic and local transport services with regard to the lines, single products or train categories if applicable train operation company, number of train changing per relation" [3] 43 http://dx.doi.org/10.14311/app.2017.11.0043 http://ojs.cvut.cz/ojs/index.php/app z. michl, m. drábek, r. vávra acta polytechnica ctu proceedings sma stresses that "a single line represents each service (and the corresponding journey in the reverse direction) connecting the stations it passes through on its trip, with the arrival and departure times for each station indicated. the repeating aspect of the timetable means that a whole repeating family of trains can be covered by one line on the diagram. the connection times between services at a station can be seen at a glance, giving a powerful method for improving the coordination between services." [4] the notation rules represented below generally come out from both above mentioned netgraph providers. besides fbs/iplan and viriato, some other timetabling software tools, e.g. railsys [5], optitakt [6] and takt [7] can produce netgraphs as well. 2.1. nodes nodes of a netgraph represent important places of a transport network from the passenger’s (user’s) point of view. each node must be named (in full or using an abbreviation – for the recommended place for descriptor see figure 1). optionally, a minimum transfer (connection) time in a node in minutes may be added to the description. node importance category, frequency of service and other attributes may be displayed using different borderline and filling types and colours. a netgraph node depicts a timetable and network node, which is: • station of origin or destination of at least one public transport line (or a part of its services), • transfer (interchange) station to other public transport lines and transport systems, even if they are not included in the netgraph • stations, where splitting or joining of trains or lines or other technology operation regularly takes place, • other important places in the network, such as regional or tariff borders, crossing of trains etc., according to the purpose of creation of the netgraph. 2.2. edges a netgraph edge represents a public transport line (or its branches and joint sections) from the passenger’s point of view. edges of lines passing more nodes are connected through the node boxes. important edge attributes are line style and colour, which depicts the service function and interval of the public transport line. 2.3. edge colour colours are usually used (according to the purpose of creation of the netgraph) for: • transport segment of the line (e.g. train category), • operation periods (e.g. lines operating only on working days, during peak hours etc.). figure 1. a node of a netgraph including name descriptor and the minimum transfer time (5 min). • different transport modes (each mode in another colour), • different train operating companies, public transport authorities etc., where applicable, • in exceptional cases individual lines can be distinguished by the line colour. there is no single correct way to colorize the lines. the selected scheme therefore has always to be explained in the legend next to the netgraph. yet there are common conventions. the authors recommend getting inspired from corresponding use cases. 2.4. edge line style a line style of the edge distinguishes the interval between services of the line. line style can even show public transport lines operated without a fixed interval. generally, one solid line represents one service per period (usually one hour). by convention, the line styles as shown in figure 2 are used. different line styles are used e.g. in urban public transport systems. in each case, the line styles used in a netgraph should always be explained in the legend. 2.5. timetable notation edges representing public transport lines with periodic timetable are complemented with time labels. they indicate periodic arrival and departure times to and from each node, noted only in minutes. the descriptors are usually on the side of travel direction from the edge – on the right in most cases, on the left e.g. for swiss railways. the descriptor closer to a node shows the arrival time, the one further from the node then the departure time respectively, as shown in figure 3. both lines (black and red) operate every hour between a and c. trains of the black line leave a every hour at minute 31 for c. they arrive to b in minute 44 and leave in minute 45, while arriving to c in minute 59. in the opposite direction the trains leave c every hour at minute 1, call at b between minutes 15 and 16 and reach a in minute 29. trains of the red line leaves a every hour at minute 4, does not call at b and arrive to c at minute 27. in the opposite direction the trains of the red line leave c every hour in minute 33, do not call at b and reach a in minute 56. if trains of a public transport line 44 vol. 11/2017 netgraph figure 2. line style usage convention according to the service interval. do not call at the node, the time descriptor is usually omitted also (see the red line at node b in figure 3). in some cases of netgraphs for timetabling or railway capacity planning purpose, it may also be useful to indicate the passage time as departure times. it is then useful to distinguish the nodes where the train calls with dots in the colour of the respective line at intersections of the edge and node borders (see figure 4). this is also useful for lines with fixed stops but non-periodic timetable, where minute descriptors are omitted. in the case of two-hours-interval the descriptor must also indicate the parity of the hour of operation (even or odd). by convention, the minute-descriptor of an even hour is descripted by normal font and that of an odd hour in italics, see figure 5. the public transport line is operated every two hours between a and c. for c it leaves a every even hour at minute 40, arrives at b every odd hour in minute 14, calls until minute 16 and arrives at c every odd hour in minute 59. in the opposite direction, it leaves c every even hour at minute 01, calls at b every even hour in minutes 44 to 46 and reaches a every odd hour in minute 20. alternative resolution of even and odd hours is also possible, e.g. by underscoring the minute descriptor for an odd hour. in the case of public transport lines operated every timetable period (60 minutes), there is no need to distinguish between odd and even hours, see both lines in figure 4. for public transport lines operated more than once a timetable period (with notation as in figure 2), only one minute-descriptor for all trains of the line within a period is provided, other service times of the period are omitted – see figure 6. the public transport line is operated twice an hour (every 30 minutes) between nodes a and c. for c it leaves a every hour in minute 03 and 33, arrives at b every hour in minute 29 and 59, calls there until minute 30 and 00 and arrives at c every hour in minute 58 and 28. in the opposite direction, it leaves c every hour in minute 02 and 32, calls at b every hour in minutes 30 and 00 to 31 and 01 and reaches a every hour in minute 57 and 27. descriptors for the opposite direction are calculated symmetrically along the main timetable axis (as complements to 60 or another timetable period). 2.6. number of serviced stops intermediate stops of a public transport line are shown on the edge between nodes. either each serviced stop is indicated with a solid dot (see figure 7) or grouped in an empty circle with number of grouped stops next to it or in it (see figures 8 and 9). 2.7. branching of a public transport line branching (some trains continue from a node in other direction than the others) is indicated in the node with two connections between the edges, see figure 10. trains arrive from b to a every hour in minute 29. they continue every odd hour in minute 32 for d, and every even hour in minute 32 for c. the situation is similar in the opposite direction. the line style shows the doubled service interval of the branches. 2.8. splitting and joining of trains (units) in this case, each train splits into two parts, and each part continues to a different destination, see figure 11. trains from b arrive in a every hour in minute 28 and split in two parts, one leaving for d every hour in minute 32, the other one continuing to c every hour in minute 35. the situation is similar in the opposite direction. the form of linking of edges in a node can also indicate a u-turn of trains of the public transport line. 3. proposed additional rules for netgraph creation and discussion thereof 3.1. different service routing in each direction in some cases, a route of a public transport line contains sections, where services operate in one direction only. such sections are distinguished with arrows. time descriptors are displayed in particular direction only. in figure 12, the line is from a to f routed through nodes b, c and e, and in the opposite direction through nodes e, d and b. 3.2. alternative routing of freight train paths lindner and von redern [8] found that it was necessary to let periodic time windows ("canals") free for freight 45 z. michl, m. drábek, r. vávra acta polytechnica ctu proceedings figure 3. example of timetable descriptors. figure 4. extended timetable descriptors example. serviced nodes are distinguished with dots of respective colour of the line. figure 5. extended timetable descriptors example – 2h interval with odd hour distinguished. figure 6. extended timetable descriptors example – 30 min interval. figure 7. three serviced intermediate stops between nodes a and b, each one indicated separately with a solid dot. 46 vol. 11/2017 netgraph figure 8. five intermediate stops between nodes a and b, the number of grouped stops is indicated with a numeric descriptor next to an empty circle. figure 9. five intermediate stops between nodes a and b, the number of grouped stops is indicated with a numeric descriptor in an empty circle. trains. these should be wide enough for required number of freight train paths per hour. these time windows should be preferably connected in nodes (if needed, among more than two railway lines to enable use by freight trains with various origin/destination stations). in the case of insufficient capacity, the authors proposed reviewing the structure of passenger transport offer. this idea was further worked up by drábek [9], who proposed a concept of network-bound periodic freight train paths with possible alternative routing in node station. this concept is, to a certain extent, an analogy to integrated periodic timetable (ipt) in passenger transport. ipt is a network-wide, periodic offer of services and transfer connections between them. the concept of network-bound periodic freight train paths is a network-wide periodic offer of capacity that meets various requirements of freight train operating companies. unified period and unified (zero) symmetry are maintained in this concept as well. alternative routing through nodes represents an analogy to transfer connection in ipt. this alternative routing is illustrated in figure 13. the train paths from b through a are routed alternatively in the direction of c and d. a freight train can pass through node a (without stop) in minute 32, either in direction c or d. in the opposite direction, a freight train can pass through node a in minute 31, either from direction c or d. 3.3. train crossing on a single-track line for planning of periodic capacity of single-track lines, a key information is in which stations the trains in opposite directions periodically cross (meet) each other – at present day, or in a certain planning time horizon. in the case of ipt, a unified axis (minute) of symmetry is kept, and the crossing happens (at least approximately) in symmetry time. in europe, a convention on zero symmetry is kept. due to mathematical prefigure 10. branching of a public transport line into two directions. figure 11. splitting and joining of trains (units). conditions of ipt, two groups of nodes with symmetry time always occur – for instance, for 60 min interval, zero symmetry means one group of nodes (stations) with symmetry time in minute 00, and another group of nodes with symmetry time in minute 30. necessity of displaying time descriptors in such symmetrical crossing stations depends on level of detail for the capacity planning. for timetabling purposes, the descriptors are necessary for information about (real or required) scheduled running time between neighbouring symmetry nodes. for conceptual planning, name of the node station may be sufficient, as such symmetrical crossing stations change only after implementation of a new ipt-stage (shortening of travel times). figure 14 shows, beyond a "common" netgraph, depiction of periodic crossing of fast train in station katovice. as this crossing is not connected with a commercial stop and the netgraph serves for conceptual level of detail only, time descriptors for katovice are omitted. 3.4. discussion besides timetable concept elaboration and presentation, netgraphs can be used for railway capacity planning as well. the authors further propose additional rules to address various aspects of railway capacity planning. in many cases, the meaning of an edge, displayed in a netgraph, shifts from periodic service to periodic capacity, which may (fully or partially), or may not, be really used by a train (service). 47 z. michl, m. drábek, r. vávra acta polytechnica ctu proceedings figure 12. operation of a public transport line in a certain direction only. figure 13. alternative routing of freight trains (alternative train paths – from a either to c or d). figure 14. example of descriptor of a crossing station (zero symmetry) [10]. 4. conclusions a netgraph is a very useful tool not only for public transport (mostly railway) timetable conceptual planning. for periodic railway capacity planning (either periodic passenger service with half interval during the peak hours, or periodic capacity for freight trains, even with alternative routing). netgraph cannot replace classical timetabling tools, but can enable an easy and rapid overview for network capacity planning. for codification of necessary descriptors of symmetrical periodic crossing in terms of group of symmetry nodes, all-day or peak only incidence and planning horizon, further research is needed. references [1] irfp e.k. institut für regional-und fernverkehrsplanung. our product: the timetable construction system fbs [online]. [cit. 2017-03-24] "http://irfp.de/our-product-fbs.html". [2] s. und partner ag: viriato, zlr. the versatile timetable system [online]. [cit. 2017-03-23] "http://www.sma-partner.com/en/solutions/viriatoand-zlr/overview". [3] irfp e.k. institut für regional-und fernverkehrsplanung. timetable creation [online]. [cit. 2017-03-23] "http://irfp.de/timetable-creation.html". [4] s. und partner ag: viriato, zlr. netgraph. focussing on service [online]. [cit. 2017-03-23] "http://www.sma-partner.com/en/solutions/viriatoand-zlr/functionality/netgraph". [5] r. m. c. gmbh. railsys [online]. [cit. 2017-03-17] "http://www.rmcon.de/railsys-en/". [6] w. h. et al. optitakt [online]. [cit. 2017-03-24] "http://www.optitakt.de/". [7] c. o. t. f. s. u dresden. timetabling [online]. [cit. 2017-03-23] "https://tudresden.de/bu/verkehr/ila/vkstrl/forschung/ forschungsfelder/fahrplanung?set_language=en". [8] h.-r. lindner, h. von redern. güterzüge im taktfahrplan – möglichkeiten und grenzen. die bundesbahn 10/1989. [9] m. drábek. periodic freight train paths in network. doctoral thesis, 2014. "http://takt.fd.cvut.cz/cargo/drabek_thesis.pdf". [10] m. drábek. netgraphs of regional railway public service in chosen czech regions. ctu in prague, undated. unpublished. 48 acta polytechnica ctu proceedings 11:43–48, 2017 1 introduction 2 common practice in netgraph notation 2.1 nodes 2.2 edges 2.3 edge colour 2.4 edge line style 2.5 timetable notation 2.6 number of serviced stops 2.7 branching of a public transport line 2.8 splitting and joining of trains (units) 3 proposed additional rules for netgraph creation and discussion thereof 3.1 different service routing in each direction 3.2 alternative routing of freight train paths 3.3 train crossing on a single-track line 3.4 discussion 4 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.10.0043 acta polytechnica ctu proceedings 10:43–47, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app modelling rock mass improvement using rock bolts jan pruška czech technical university in prague, faculty of civil engineering, thákurova 7/2077, prague 6, czech republic correspondence: pruska@fsv.cvut.cz abstract. anchor support currently represents a significant reinforcing technique in underground constructions. the principles of rock bolt reinforcement action are derived from various concepts of the underground excavation stability. in recent times, rock bolt design techniques have been complemented by numerical calculations procedures. the paper describes fem modelling of a rock bolt system for mechanical and grouted bolts. keywords: rock bolts, modelling, fem. 1. introduction in recent times, rock bolting represents a significant reinforcing method, especially in tunnelling. the main effect of rock bolting is in reducing and eliminating the impact of different types of failures that occur in the rock mass (schistosity, joints, cracks, etc.) so that the reinforced rock mass can restore the equilibrium disturbed by excavating an underground work. the effect of rock bolting on the rock mass is derived using a variety of methods to determine the stability of underground excavations and the rock bolt type considered (e.g. mechanical or resin). the design and verification of the structural behaviour of rock bolts can be made by any of the following methods: • empirical methods or using rock mass classification, • analytical solution (e.g. rock reinforcement contact problem), • simple calculations based on the fenner-pacher curve theory, • mathematical modelling (numerical solution of a contact problem), • models based on discontinuity mechanics theory, • probabilistic (stochastic) approach to determine the required loading of reinforcement anchorage. of the above methods, mathematical (numerical) modelling using the finite element method (fem) is commonly used in todayâăźs practice. generally, the design of bolts is a complex task that requires the knowledge of rock bolt materials, the anchor technology and the rock mass behaviour. 2. rock bolts modelling using fem the modelling of anchors and rock bolts by fem is more complicated compared with the modelling of the rock mass itself. this is mainly caused by the fact that rock bolts are divided by the type of anchor (simply into two main groups: point-anchored and anchored along its entire length or predominantly over the entire length). thus, the fem model must necessarily take into consideration the type of rock bolt, the anchoring method and the distribution of anchoring forces in the surrounding rock. it can be used for the following task: • modelling a rock bolt as a single element, • homogenization of the rock bolt’s area. in current tunnelling practice, rock bolts attached to the rock along their entire length predominate, thus, the methods using the homogenization of the rock bolt’s area are mostly used. 2.1. modelling of rock bolts using homogenization the principle of this procedure is based on the fact that anchors create a supporting ring in the rock mass near the excavation with characteristics of a thickwalled cylinder. the area of the supporting ring is modelled so that the elements of the finite element mesh in the supporting ring change their geomechanical properties by selecting equivalent isotropic material with improved cohesion. however, this procedure is not suitable for a clearly discontinuous rock mass (connection of rock blocks using rock bolts) because it does not consider the mechanical strength of the rock bolt’s area. czech practice uses the following procedure to increase cohesion in the rock bolt’s area. the increase of cohesion due to bolting is calculated according to the equation: cs = nu + sin ϕ′ ak · 2 · cos ϕ′ · γkc (1) where: nu loading capacity of the bolt [kn], ak area per rock bolt [m2], ϕ′ effective angle of internal friction [°], γkc coefficient of bolting reliability [−]. 43 http://dx.doi.org/10.14311/app.2017.10.0043 http://ojs.cvut.cz/ojs/index.php/app jan pruška acta polytechnica ctu proceedings cohesion of equivalent isotropic material: ch+s = ch + cs (2) where: ch+s total cohesion of rock increased due to rock bolting, ch original cohesion of rock, cs cohesion increased due to rock bolting. if we need to take into the account the impact of the bolt on the discontinuity surface, a so-called virtual element can be used in fem. this virtual element models discontinuity surfaces using a conventional rectangular element with a height equal to the length of the plastic area in the bolt (fig. 1). it is assumed for the shear modulus calculation that a failure will occur in the bolt and at the discontinuity interface at the same displacement. the shear modulus was expressed e.g. by h. larsson [1]: gv p = τb uf · e · ab at + τj uf · e · aj at (3) where: gv p shear modulus of the virtual element, τb shear stress at the rock bolt failure, uf block displacement at the discontinuity interface, e height of the virtual element (length of the plastic area in the bolt), ab rock bolt area, aj discontinuity interface, τj shear stress at the discontinuity interface at failure, at total area of the rock with rock bolts (ab+aj). 2.2. modelling of rock bolts as discrete elements the problem of the formulation of a finite element that would describe the stabilizing effect of the rock bolt or rock anchors on the rock mass has been solved in various approaches since early fem numerical analyses (the first models were derived around 1968). thus, a number of special elements for describing rock bolts were developed (for bolts fixed at both ends or along their own length). these special elements are connected with specific special software or they require direct implementation in a programme code by the user. in the following part, one of the most used elements (attached to the rock along their own length) will be described. the principle of this element was presented in the thesis by ömer aydan in 1988 [2]. it is essentially a beam element (special bolt element), which has four nodes, two of which are connected to the anchor rod and the other two to the surrounding rock fig. 2. the following several simplifications have been considered for deriving the stiffness matrix of this element: • the body formed by the grouting material is considered as an axisymmetric thick-walled hollow cylinder, which, due to its shear deformation, transmits tension from the anchor rod into the surrounding rock. • both bolt materials (grouting material and the anchor rod) are considered with a linear elastic behaviour, and thus the relation between the stress and strain in the anchor rod and in the grouting material are described using hooke’s law. • in terms of the coordinates of nodes, the radius of the bolt is negligible in terms of the coordinates of nodes. this presumption naturally does not apply when creating the stiffness matrix. • the bolt considers only the existence of three types of deformation: . ε0 relative longitudinal deformation of the anchor rod caused by axial displacements of the nodes of the anchor rod, . γg shear deformation of grouting caused by different axial displacements of the nodes of anchor rods and the nodes connected with the surrounding rock, . γx transverse shear deformation of grouting caused by different radial displacements of the nodes connected with the surrounding rock. the relationship of the rod bolt strain in terms of nodal displacements (the b index determines the variables associated with the bolt (rod) and the g index determines the values belonging to cement grout): εb = bb ub (4) and the relationship of the grouting ring strain using nodal displacements εg = bg ∆ug (5) where u is the matrix of displacement and b the matrix of differential operators. physical equations for the linear elastic behaviour of rods and the grouting ring can be described using the above expressions as: σb = db εb σg = dg εg (6) where d is the relevant material stiffness matrix. by discretization using finite elements the equations of virtual work (it is assumed that the stress field σ is structurally possible and the field of displacements is kinematically allowed), we get expressions for the stiffness matrix: ke = ∫ b t db d(vol) = ∫ le 0 ∫ rh 0 ∫ 2π 0 b t db dϕdrdz (7) after the integration, we obtain the stiffness matrix of the element in the form given in fig. 3, where 44 vol. 10/2017 modelling rock mass improvement using rock bolts figure 1. virtual element for the rock with bolts. figure 2. scheme of a special bolt element: 1, 2 nodes connected with surrounding rock; 3, 4 nodes connected with the bolt rod. kb = eb · π · r2b le (8) ks = gb · π · r2b le (9) kg = π · gg · le 3ln( rh rb ) (10) aydan’s element has been successfully implemented in scientific and commercial programmes. fig. 4 shows an example of the application of this element to the pmd system by d. runt [3]. 2.3. modelling using geotextile elements some authors in their publications describe the ability to model grouting bolts using geotextile elements e.g. m. šejnoha [4]. these are reinforcing elements which are continuously fixed in the earth massif. this element is, therefore, a system of two node rod elements connected in the places where the line of the geotextile intersects the edge of individual network elements see fig. 5. these elements can only be loaded by tensile forces, they do not have compressive strength and cannot transmit bending moments. they exhibit linear elastic behaviour (that is dependent on a specified compressive strength) and have limited tensile strength until the moment of failure. 2.4. modelling using internal forces in the rock mass modelling using rod elements is another possibility for describing bolts that transmit only axial forces. these elements have the ability to transfer compressive load, unlike the elements for geotextile modelling. it is also possible for this element to specify the input forces in the anchor (prestress) and not consider the prestress due to compressive load. according to some authors (e.g. [5]), the bolt root can be modelled as elements without compressive strength with contact elements. the cohesion of contact elements should be determined so that the total frictional force acting on a 1 m strip corresponds to the frictional force acting on the bolt (reducing cohesion by the ratio of the bolt area to the finite element area). for modelling bolts fixed along their own length, this model is unacceptable in underground engineering [6]. 2.5. modelling using internal forces in the rock mass the use of internal forces is one of the oldest methods of modelling anchors and rock bolts in fem. the principle of modelling consists in the insertion of internal forces into the rock mass at both sides of the rock bolt (head and root). this kind of plane modelling is only used for tensioned bolts or anchors. the inserted forces simulate the bracing area in the rock mass stabilized by rock bolts. however, this kind of modelling often leads to the plastification in the internal force location and to a situation where the calculation does not converge. 3. effect of installation time it is s quite difficult to make a sufficiently accurate model of the effect of rock bolts. we can usually only estimate the maximum possible effect. this is not only due to the established simplifications in different procedures, but the installation time of rock bolts also plays an important role (time delay due to the 45 jan pruška acta polytechnica ctu proceedings figure 3. the stiffness matrix of the element figure 4. rock behaviour in the stress σxx curve without rock bolts (left) and with rock bolts (right) figure 5. dividing geotextiles on rod elements application of shotcrete lining). this effect is very difficult to describe correctly in a numerical model. rock bolts are often part of primary lining and also help to improve the final deformation. therefore, we can only determine correctly the effects of the immediate installation of bolts or bolts installed to partially deformed lining. 4. conclusion this paper gives an overview of different approaches to rock bolts modelling using fem for tasks related to underground structures. from the above mentioned methods, a survey clearly shows that the main approximation of rock bolts in fem uses modified strength and deformation parameters in the rock bolt’s area or specially derived elements (quantifying the effect of bolting for the rock mass). also, fem is not suitable for modelling a discontinuous rock mass and, therefore, a direct impact of rock bolts on discontinuities; we can say that both of the above mentioned approaches (based on the theory of continuum mechanics) are quite suitable for describing the behaviour of rock bolts. subsequently, engineers can choose such bolt properties to reach a steady state of an underground structure and also optimize its design (based on the comparison of fem modelling of different options). 5. acknowledgement this article was written as part of a study within the project supported by the competence centres program of the technology agency of the czech republic (ta cr), no. te01020168. references [1] h. larsson, t. olofsson. bolt action in jointed rock. in rock bolting : theory and application in mining and underground construction : proceedings of the international symposium on rock bolting, abisko, 28 august 2 september 1983, pp. 33–46. balkema publishers, a.a. / taylor & francis the netherlands, 1984. godkänd; 1984; 20080926 (ysko). [2] ö. aydan. the stabilisation of rock engineering structures by rockbolts. nagoya: nagoya university, 1989. 46 vol. 10/2017 modelling rock mass improvement using rock bolts [3] d. runt, j. pruška, j. novotný. modelling of rock bolts using the finite element method. in isrm regional symposium eurock 2015. international society for rock mechanics, 2015. [4] m. šejnoha. finite element analysis in geotechnical design. saxe-coburg publications, kippen, stirling, scotland, 2015. [5] z. bittnar, j. šejnoha. numerické metody mechaniky 1. české vysoké učení technické v praze, 1992. [6] m. hilar. numerická analýza tektonicky porušeného horninového masivu s primární výstrojí při aplikaci nrtm. ph.d. thesis, české vysoké učení technické v praze, 2000. 47 acta polytechnica ctu proceedings 10:44–48, 2017 1 introduction 2 rock bolts modelling using fem 2.1 modelling of rock bolts using homogenization 2.2 modelling of rock bolts as discrete elements 2.3 modelling using geotextile elements 2.4 modelling using internal forces in the rock mass 2.5 modelling using internal forces in the rock mass 3 effect of installation time 4 conclusion 5 acknowledgement references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0051 acta polytechnica ctu proceedings 5:51–58, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app criticallity of transportation infrastructure in czech republic jan procházka∗, dana procházková czech technical university in prague, faculty of transportation sciences, praha, czech republic ∗ corresponding author: prochazka@fd.cvut.cz abstract. the paper deals with transportation infrastructure criticality because this quantity determines the state capability to overcome the critical conditions and to ensure the inhabitants survival. the criticality rates for individual types of transportation infrastructure and for the entire transportation infrastructure are determined by data from experts from the areas: transportation; transportation management in the territory; supply chains; public administration; and the integrated rescue system. the experts assessed 14 factors, which have been often used in the developed world countries, from the view of human security and development. the result values and their interpretations were determined by using the multiatribute utility theory. keywords: transportation infrastructure, system of systems, interdependences, criticality, interoperability, serviceability, human survival, state stability. 1. introduction transportation infrastructure belongs to the basic systems that make up the critical infrastructure in the european union, in developed countries and also in the czech republic. we understand the system of transportation in accordance with the current knowledge as a system of systems, i.e. the set of several overlapping systems. the most important property in such type of system is the interoperability of partial systems that depends on interdependences [1]. the interdependences are the cause of different emergent phenomena and cascade failures that cause that the transportation infrastructure does not fulfil its function, i.e. the serviceability of territory at normal, abnormal and critical conditions if it is threatened the human survival and state stability. therefore, it is necessary to manage the transportation infrastructure by the way so infrastructure criticality is acceptable [1]. the subject of this work is to assess the criticality of the transportation infrastructure in the czech republic on the basis of the integral safety concept, which aims are security and development of the territory concerned and its inhabitants. the concept used is based on the knowledge of the reality of the dynamic development of the human system and it considers that human system conditions are normal, abnormal and critical. the capability to overcome the critical conditions, ensuring the survival of humans and to have capability to stabilize the situation we consider crucial, and therefore, we use the degrees of criticality, which are aimed at assessing the aspects associated with the given properties. for the assessment of the criticality we use criteria used in developed countries, multi-objective approach and data from experts whose qualification is certified according to the criteria used in the european union [1]. 2. current knowledge on transportation infrastructure the transportation system, which provides transportation of humans and goods, includes all modes of transportation, which, under the coordination of individual transportation systems work together and create a logistical network [2]. transportation logistics network is currently becoming a very important area of business activities, and not just in a regional and national scale, but also in international relations. the transportation network can be combined with energy as one of the foundations of the economic prosperity of states and organizations. transportation systems for their activities require a considerable amount of funds, namely in investment, operational and administrative domains, which typically create autonomous sub-sectors, the activity of which is mutually coordinated. the characteristic feature of those technological units is their capability to own separate activity, and therefore, they are often considered as closed units, which is projected to the safety concept, and at risk management it is used the simplest approach, i.e. the procedures for a closed system, in which the sources of the risk are only the technical problems and the human factor [2]. infrastructure of transportation network is currently mostly owned by the public sector; however, its development is largely determined by the requirements of private sector. simultaneously it constitutes an important element of the economy and of the citizens living way. transportation network has considerable international political influence, as it allows the international cooperation of individual states and tourism, which contributes to the integration of citizens. the transportation system is built from the separate parts that make up certain units that provide certain 51 http://dx.doi.org/10.14311/app.2016.5.0051 http://ojs.cvut.cz/ojs/index.php/app jan procházka, dana procházková acta polytechnica ctu proceedings circuit activities separately. in the serviceability of territory (creating the space for human lives, i.e. the human system), however, the units are together overlapped, i.e. they create a system of systems [2] [3] [4]. the ground of each unit is a specific type of transportation vehicles, accompanied by complex of operational and administrative procedures. the characteristic feature of these technological units is their specialized capability to separate operation, i.e. the autonomy. the capability to separate operation, however, it is only apparent, the sub units are bound by the rules of their transport logistics, which not only coordinates their activities, but also makes it economically more profitable because it reduces operating costs. on the other hand, however, it is the source of internal dependencies "interdependences" [2] [5]. safety of transportation networks in the past was provided only by a simple guarding. the guarding was mostly limited to surveillance, which should prevent the mostly property-related offences, such e.g. theft of the goods transported, pertinently endangering the personal safety of transported passengers. on the basis of current knowledge, it is the aim of the safety management of the transportation system, which consists of objects and infrastructures, ensuring the security and development of both, the transportation system and transportation system vicinity, because it is the only way how to ensure the safe human system. on the basis of current knowledge, this means that we have to apply the tools of the disciplines that provide safety management of complex systems [1] [2] [5]; i.e. the multi criterial approach based on expert assessment of the possible variants that describe the behaviour of systems of systems designated by different behaviour of individual subsystems [4]. 3. transportation system of the czech republic in the czech republic a transportation network is in essence created from the following areas: railway transport system; road transport system; air transport; and water transport. the main activities of the business in the transport are passenger and freight transport. ensuring the rational operation requires considerable resources – labour, capital, information, as in a modern industrialized society, it is common that the interconnection and operation of transportation networks and the provision of transport services is an important part of near all other activities. for example, the provision of transport services in the major cities is done by public transportation, which is based on mostly coordinated cooperation of road and rail transport, sometimes supplemented by underground transportation; it is basically so called integrated transportation system. in the last few decades, the transportation infrastructure has become a major target of terrorist attacks. in particular, it clearly reflected in the air transport, in which they terrorists originally focused only on the kidnappings, and the destruction of the aircraft in flight [5]. after the attacks of 11 september, 2001 on the trade centre in new york, it turned out that the aircraft may change in the gun with the devastating consequences of the disastrous. from historical practice it is known that natural disasters (in our condition floods, storms, earthquakes and landslides) often caused considerable and practically difficult to quantify damages to parts of the rail and road transport of technological system and the environment. therefore, it can be concluded that transportation systems are particularly vulnerable to attacks and natural disasters. they are very favourite target of the activities of terrorists who by attack on any of transportation systems can cause considerable damage to property and infrastructure devices. it is necessary to recognize that the indirect impacts of attacks and their consequences can even endanger the stability of the economics of the industry, or even the state. from the above analysis it follows that it is necessary to devote to the issue of the destruction of the infrastructure of transportation systems special attention, to define particularly sensitive parts and to search for ways how to reduce the danger of the destruction of important parts, which can cause the interruption of the activity of transportation systems, and by such way to prevent the economic destabilization not only of the relevant business units, but also to other bodies in the industrial areas. if we want to manage the transportation system safety with the aim of ensuring its security and development, so we need to know the priority aspects, on which the achievement of objectives depend and on that we need to focus the attention, i.e. the measures and activities [1] [5]. from the same source it follows that the critical spots in the technological system (object, the infrastructure, plant, facility, territory) are the places where the basic technological processes are under way and for which there are in force specific provisions for the safety under normal, abnormal and critical conditions (e.g. single-level crossing of the railways and roads in the village). table 1 gives the important parts of the transportation system, separated for four of its parts according to criteria of governmental decree no. 2000 col. [6]. 4. method for determining the transportation infrastructure criticality according to the detailed discussion given in [1] [5] the criticality of item means that item has simultaneouslygreat importance and great vulnerability. it means that we work with conflict criteria at the solution of problem. therefore, we use the strategy defined by the keeny and raiffa [7] defined by the mathematical statement "the higher, the worse, i.e. the higher criticality in our case". as it has been already said above, the transportation system as a whole and its basic subsystems are 52 vol. 5/2016 criticallity of transportation infrastructure in czech republic rail transport road transport air transport water transport linear constructions tracked rail network, the electrification network of railways, railway security network the full road network the system of interconnected airports the waterways objects the railway station buildings for traffic management, expedition of passengers and redeployment of goods, the depots of vehicles, railway workrooms, railway bridges, railway tunnels, track divisionmaintenance rail network objects and networks for security techniques, railway workrooms for servicing the locomotives of all kinds and fleet, the power stations of power supply to rail network, fuel stations road bridges, road tunnels, stocks of fuels for road transport purposes, objects for the management of road traffic, gas fuel stations for road vehicles, objects for the management of emergency situations in road network airport runways for takeoff and landing, a network of runways, lighting, passing ways, aircraft stands, buildings for service stations for air transport, buildings for normal preflight preparationinput and general examination, air traffic control objectsairport, regional, the radar station for air traffic control, radar stations for meteorology, computing centre of the airport, the equipment for operational information sharing for regional service, facilities for the sharing of operational information for the management of the activities of the airport, centre for intervention in emergency traffic incidents (accidents...), gas stations and fuel warehouse ports with the devices, the locks, docks, hydrometeorology stations, objects for maintenance of vessels, fuel stations fittings rail vehicleslocomotives and cars, equipment of the rail stations and the tracked rail network, device for controlling the places on the crossing the road and rail transport, the equipment for operational information sharing within the operation of the rail system, facilities for the supply of fuel for vehicles the equipment for operational information sharing with the road network, equipment of centres with technique for solving incidents, buses, freight cars, personal cars aircrafts ships table 1. important parts of transport system 53 jan procházka, dana procházková acta polytechnica ctu proceedings systems of systems, i.e. it goes on several mutually penetrating systems of different nature (technical, organizational, financial, human resources, and legal), and therefore, the criticality assessment cannot be performed from a single indicator, but it is necessary to use a multicriterial approach and information from experts [4]. in accordance with the procedures that are used in the usa, australia, canada etc. [8] [9] [10] [11] [12] [13] there were used the following factors important to the criticality of transportation infrastructure: • a rate of capability of defence, • a rate of vulnerability to attack, • a rate of threat to the health and lives of humans, • a rate of accidents impact on the environment • a rate of cost of exchange or repair, • a rate of exchange time or repair time, • a rate of importance for ensuring the life-saving or emergency functions in the territory, • a rate of importance for ensuring the functions of administration and local government, • a rate of importance for ensuring the functions of army and police, • a rate of redundancy or replacement services, • a rate of importance for ensuring the communication functions, • a rate of impact of frontage failure on the economy of region (state), • a rate of importance of operability, • a rate of importance in the field of symbolic, cultural, etc. with the facts in the documentation of the czech republic transportation system [14] and with the data from 5 experts, selected according to the criteria of the eu given in [1], from the areas: transportation; transportation management in the territory; supply chains; public administration; and the integrated rescue system, there wereobtained verbal assessments of the factors mentioned above. since the main human objective is the human safety and the human development [3], for which the safe transportation infrastructure is vital, the criticality of the transportation infrastructure is understood from the perspective targeted on the human survival. therefore, the verbal assessment was transferred with the help of five experts participating in the panel discussion into the numerical scale that is analogue to the scale using in the risk assessment in the čsn standards [1] – allocate number of points is the following: (1.) if factor does not contribute to the criticality of the transport system (the expected damages are less than 5%), (2.) if factor contributes little to the criticality of the transport system (damages are expected at an interval of 5-25%), (3.) if factor contributes intermediately to the criticality of the transport system (the expected damages are in the interval 25-45%), (4.) if factor contributes significantly to the criticality of the transport system (the expected damages are in the interval of 45-70%), (5.) if factor contributes very significantly to the criticality of the transport system (the expected damages are in the interval of 70-95%), (6.) if factor significantly contributes most significantly to the criticality of the transport system (the expected damage is greater than 95%). the resulting evaluation of criticality for each partial system provided that we give equal weight to all factors, can have the values 0 to 70 and for the entire system having 4 subsystems, provided that their equivalence the criticality values are between 0 and 280. if we again apply approach used in the čsn standards [1], we get the data given in table 2. 5. results – the criticality rates for transport infrastructure the verbal assessments of factors for rail transport, obtained from the experts are the following: (1.) rate of capability of defence – defence of the railway is difficult due to the considerable complexity of rail system. at the attack of terrorists on main parts or at natural disaster there may originate extensive damages that stop the operation of a large part of the system. the human factor and organizational accident can also cause considerable damages. (2.) rate of vulnerability to attack – vulnerability of the railway is significant because some parts of it can be quite simply destroyed extensively (railway stations complexes, bridges, tunnels, etc.). (3.) rate of threat to the health and lives of humans – the health and lives of passengers are threatened by practically in all the accidents of the rail system. (4.) rate of impact of railway accidents on the environment – railway accidents can have significant impacts on the environment, in particular in the transport of fuels, chemicals and other raw materials. removal of contamination of environment is the most difficult and expensive. (5.) rate of cost of exchange or repair – elimination of damages caused by the destruction of the track system can be mostly done without the big cost. expensive there are large constructions repairs of faulty bridges and tunnels. repair or restoration of the destroyed railway vehicles is often expensive. 54 vol. 5/2016 criticallity of transportation infrastructure in czech republic criticality rate scorethe percentage of the total value individual transport systems entire transport system the extremely high more than 95% more than 66.5 more than 266 very high 70 95% 49 66.5 196 266 high 45 70% 31.5 49 126 196 medium 25 45% 17.5 31.5 70 126 very small 5 25% 3.5 17.5 14 70 insignificant / negligible less than 5% less than 3.5 less than 14 table 2. value scale used for determining the criticality rate (6.) rate of exchange time or repair time – repair of tracks is quite fast. large construction repairs of bridges and tunnels require a longer period of time. (7.) rate of importance for ensuring the life-saving or emergency functions in the territory – in emergency situations the railway is very important and often essential for the provision of emergency activities. (8.) rate of importance for ensuring the functions of administration and local government – for the provision of administrative and self-government the railway is important, but its services for the activities in most cases can be replaced by other means of transport. (9.) rate of importance for ensuring the functions of army and police – to the needs of the army and the police the railroad is necessary. (10.) rate of redundancy or replacement services – backup at the railway is quite high (multiple lines), particularly in the transport of basic industrial raw materials, food, and fuel in larger quantities. very important it is container shipping service linking the various transport systems and container transport requires its own technology. in particular, the important role has container transport in international traffic, which of course, requires the construction of investment-intensive terminals, which can be a sensitive target for terrorists. (11.) rate of importance for ensuring the communication functions – in emergency case, the communication network of railway can be used because it is independent. (12.) rate of impact of frontage failure on the economy of region (state) – disturbed rail traffic system adversely affects the smooth running of the regional or national economy and causes considerable economic losses. (13.) rate of importance of operability – railway operability is very important on both, the national and the international scales. (14.) rate of importance in the field of symbolic, cultural, etc. – .-the level and status of the railway transport system is testimony to the cultural and technical level of the state. the verbal assessments of factors for road transport, obtained from the experts are the following: (1.) rate of capability of defence – protection of the road network is complex because even if the disruption of the custom paths can be quite easily and quickly repair, sensitive parts such as bridges and tunnels it is necessary against terrorist attacks and natural disasters especially consistently to protect, because their damage or destruction often requires a long-term repair, mostly investment-intensive. the human factor and organizational accident can also cause considerable damages. (2.) rate of vulnerability to attack – vulnerability of road network is particularly determined by sensitive parts of system, such as bridges, tunnels, etc. (3.) rate of threat to the health and lives of humans health and lives are threatened by accidents during the operation, the number of which can be in the effective management of road traffic minimised. (4.) rate of impact of road accidents on the environment road transport heavily polluting environment by gases that vehicles emit into the atmosphere. when accidents of vehicles, environmental pollution is caused in particular by spilling fuel and possibly also by transported goods that is spread in nature. (5.) rate of cost of exchange or repair – except the costly repairs of sensitive network places – bridges tunnels etc. it is possible to perform repairs of damages of road network fairly cheaply and in a short time. (6.) rate of exchange time or repair time – outside the repairs of bridges, tunnels and other expensive parts it is possible to perform the repairs of network in a relatively short time. (7.) rate of importance for ensuring the life-saving or emergency functions in the territory – the serviceability of roads is very important, in the case of an emergency it can be solved by the finding a replacement path that avoids the damaged parts. (8.) rate of importance for ensuring the functions of administration and local government – road transport is very necessary. 55 jan procházka, dana procházková acta polytechnica ctu proceedings (9.) rate of importance for ensuring the functions of army and police – road transport is fundamentally necessary. (10.) rate of redundancy or replacement services – there are, in principle, to allow mutual logistic redundancy compensation of individual transport systems. (11.) rate of importance for ensuring the communication functions – communication function can be partially replaced by road transport. (12.) rate of impact of frontage failure on the economy of region (state) – the status and level of road system quality has a significant influence on the economy. (13.) rate of importance of operability – operability of the road system is very important and fundamentally affects all activities in the region and state. (14.) rate of importance in the field of symbolic, cultural, etc. – the level of road transport and the road system is indicative of the level of the state and its tradition. the verbal assessments of factors for air transport, obtained from the experts are the following: (1.) rate of capability of defence – defence of air transport is very exacting and difficult, because the air transport system is currently the most complex of all transport systems. the accidents in aviation occur from reasons, the technical reasons and the meteorological reasons. however, the air transport is also a hot target for terrorists. the human factor and organizational accident can also cause considerable damages. (2.) rate of vulnerability to attack – the vulnerability is high. preparation of the aircraft for operation is very challenging. therefore, the flight and technical staff is carefully selected and trained. defence against terrorist attacks is performed while the intensely, but not always successfully. (3.) rate of threat to the health and lives of humans – in air accidents it is occurred injury and death of crew and passengers. therefore, the prevention of accidents in the aviation service is very important. (4.) rate of impact of road accidents on the environment – air accidents are almost always accompanied by considerable local damage to the environment, but mostly they are hitting a small part of the territory. (5.) rate of cost of exchange or repair – replacement of the damaged aircraft technique is almost always very costly financially and time often challenging. repairs require a considerable supply of spare parts and the repair work of specialists. the repair needs to be carried out at the manufacturers or authorised service companies. (6.) rate of exchange time or repair time – the time length of the repair is different according to the nature of the fault. (7.) rate of importance for ensuring the life – saving or emergency functions in the territory – aviation equipment has a significant position and it is almost necessary in response to real crisis situations of all kinds. (8.) rate of importance for ensuring the functions of administration and local government – it is important only conditionally. (9.) rate of importance for ensuring the functions of army and police – it is very necessary. (10.) rate of redundancy or replacement services – air transport has a conditional redundancy that from its independence on surface networks enables to solve specific tasks. (11.) rate of importance for ensuring the communication functions – in some exceptional cases, air transport is the only viable solution. (12.) rate of impact of frontage failure on the economy of region (state) – air transport is of great importance for the economy, because outside of its own air traffic it is driving article of industry. (13.) rate of importance of operability – it is very important. (14.) rate of importance in the field of symbolic, cultural, etc. – air transport in the czech republic has a significant position in the world. the aviation industry of the czech republic has a long tradition, the products have been exported into all world. the verbal assessments of factors for water transport, obtained from the experts are the following: (1.) rate of capability of defence – with regard to the limited scope of the network of waterways in the czech republic this transport system can be probably successfully protected. the human factor and organizational accident can also cause damages. (2.) rate of vulnerability to attack – the vulnerability of the network of waterways, in particular, determines the sensitive parts – locks, weirs, etc., that ensure the navigability of the waterways. (3.) rate of threat to the health and lives of humans – the health and lives of the residents and workers in the transport sector are threatened to the extent that is usual in other sectors of industry. (4.) rate of impact of road accidents on the environment – in particular the pollution of waterways by fuels, which adversely affects the environment and the biosphere in the water and on the banks of streams. (5.) rate of cost of exchange or repair – ship accident are not common, there is not usually a large extent of the damages. 56 vol. 5/2016 criticallity of transportation infrastructure in czech republic factor rail transport road transport air transport water transport entire transport 1 5 4 5 2 16 2 5 4 5 3 17 3 4 3 4 3 14 4 4 3 3 2 12 5 4 3 5 2 14 6 3 2 4 3 12 7 5 5 4 1 15 8 3 3 2 2 10 9 5 5 5 2 17 10 3 2 3 1 9 11 3 2 3 1 9 12 5 4 4 4 17 13 4 4 4 2 14 14 3 2 3 2 10 all factors 56 46 54 30 186 table 3. rates of criticality of the transportation infrastructure in the czech republic (6.) rate of exchange time or repair time – the ship drives are mostly repairs in the port, where the repair can be done in a short time. repair of damage to the ship’s body can usually only be performed in the dock, which prolongs the time of repair. (7.) rate of importance for ensuring the life-saving or emergency functions in the territory – it is needed only for the rescue work of a certain character. (8.) rate of importance for ensuring the functions of administration and local government – small. (9.) rate of importance for ensuring the functions of army and police – it is significant only in special cases. (10.) rate of redundancy or replacement services – water transport can be replaced by the rail or road transport. (11.) rate of importance for ensuring the communication functions – small. (12.) rate of impact of frontage failure on the economy of region (state) – water transport by its low expenses positively affects the transport of raw materials and objects of large dimensions. (13.) rate of importance of operability – smaller, only to a limited extent. (14.) rate of importance in the field of symbolic, cultural, etc. – water transport has probably the longest tradition and by this reason it is one of the symbols of the country. the criticality rates determined by the methodological procedure described above are in table 3. table 3 shows that the greatest criticality has rail transport, which is followed by air transport, road transport and water transport. according to the chosen concept, the chosen scale and evaluation provided by the experts the criticality rates of both, the rail transport and the air transport are very high, the criticality rate of road traffic is high and the criticality rate of water transport is medium, and the criticality rate of whole transport system is high. from table 3 it is clear that to the size of criticality rate mostly contributes the vulnerability to attacks, little capability of protection (the protection of long linear networks in the territory is always a big problem), the importance for ensuring the functions of the army and the police, a big impact of transport failure on the economy of region (state), and the big importance for ensuring the emergency and rescue functions in the territory. 6. proposals for transport infrastructure safety improve from the above analysis of transportation systems criticality it follows that the transport domain is one of the most important areas of economic and social item in the state. to its high criticality it contributes that in recent decades this fact was discovered by population groups in different states, which are not satisfied with the current state establishment in the home countries, they create their own political ideology, which most strongly criticizes the current political establishment of the state in which they live and they try to their views and goals in life violently to enforce by illegal way and as onerous procedure for the enforcement of their ideology, they consider terrorism. terrorism currently has a global form, so its international character leads to the need for protection against attacks. in the reference case, the facts: the great importance of transport systems, the high vulnerability of transport systems, and also the fact that part of the transportation systems can be used as an offensive weapon with great destructive potential, as shown by the attacks of 11 september 2001 in new york. it is, therefore, necessary to develop countermeasures for reducing the eventuality of attack on transportation systems and to avert the devastating impacts on the functioning the state and on the lives of citizens. in conjunction with the results of the project the focus of the european union [15], it is necessary to consider that not only terrorism, but also a natural disasters, organizational accidents and corruption associated with the areas of control often to cause the extensive damages on the transportation systems. preparation and production of protective measures requires substantial investment, and therefore, it is necessary to coordinate investment and to combine possible measures for protection of transportation systems against possible damages caused by disasters [16] and terrorist activities, as required by the principles of strategic management [3]. 57 jan procházka, dana procházková acta polytechnica ctu proceedings due to the fact that neither terrorism nor natural disasters cannot be averted, the main goal of our activities are mitigating measures at occurrence of disasters in order to reduce the losses of human lives and minimize damages [1] [3]. because of the importance of transportation systems is considerable, it is very urgently to implement the measures and activities that will allow the critical infrastructure elements of transportation systems effectively protect against the activities of international terrorism and other disasters. it is necessary to carry out measures that reduce the vulnerability, ensure a rapid response to the failure of critical elements, and ensure the continuity of critical elements, and on the quick recovery of the other important elements. it is necessary to: • to estimate the vulnerability of functional elements of transportation system in the context of the assessment of transportation system objects such as bridges, tunnels, roads, highways, airports and waterways, • to provide for the most vulnerable part of the infrastructures, the failure of which can cause the most economic damages in infrastructure up to make impossible the transport activity, • to create scenarios of the impacts of major disasters and possible attacks, and on the basis of scenarios of impacts to determine the rapid-response procedures and the procedures for ensuring the transportation services as much as possible, • to improve measures and activities in the areas of planning, with the aim of improving the protection against possible disasters and terrorist attacks on the transportation system, and not only from the point of view of individual systems, but in the logistics context to link measures and to increase the safety of the entire system taken in as a system of systems, • to provide for measures that reduce the option of attack on transportation infrastructures and improve the detection of attacks on transportation infrastructure, • to estimate the amount of investment and labour costs, which will require the implementation of mitigation measures in case of failure of transportation infrastructures and to propose a procedure for their application, • to perform the analysis and to design the transportation infrastructure network protection against damages, which may arise as a result of possible disasters and to perform the estimation of operational inability origin as a result of their putting out of operation of functional elements or their parts, • to reduce the frequency of interruptions of transportation systems operation as a result of a failure that cannot be reliably in advance revealed during the regular inspections. 7. conclusions the above facts show that it is urgent to pay increased attention to safety issues of transportation systems. it needs to be processed as a separate concept of transportation system safety, so the concept of critical infrastructure safety, as it shows the work [17]. from the professional point of view, it is necessary to include into safety issues solution a systemic approach, which enables us to support the technological development of transportation systems, to consider the needs of people and transportation serviceability, the environmental protection and to reduce economic damages. it is necessary to define by the expert way the issues that are necessary to solve, and to process the relevant studies in the meaning of the above recommended procedure. references [1] d. procházková. základy řízení bezpečnosti kritické infrastruktury, 2013. isbn 978-80-01-05245-7. [2] d. procházková. bezpečnost kritické infrastruktury, 2012. isbn 978-80-01-05103-0. [3] d. procházková. strategické řízení bezpečnosti území a organizace, 2011. isbn 978-80-01-04844-3. [4] d. procházková. analýza a řízení rizik, 2011. isbn 978-80-01-04841-2. [5] d. procházková. safety of complex technological facilities, 2015. isbn 978-3-659-74632-1, id: 200303845. [6] czech republic, czech code. [7] r. l. keeny, h. raiffa. decision analysis with multiple conflicting objectives, 1976. [8] ema. critical infrastructure emergency risk management and assurance [online], 2003. www.ema.gov.au. [9] fema. promoting critical infrastructure protection by emergency managers and first responders [online], 2005. www.usfa.fema.gov. [10] saic. a guide to highway vulnerability assessment for critical asset identification and protection, 2002. national cooperative highway research program project 20-07/task 151b. [11] s. (sandia national laboratories). vital area identification, 2005. albuquerque. [12] us. the national strategy for the physical protection of critical infrastructures and key assets, 2000. [13] us. guide for critical infrastructure protection, 2005. [14] ministerstvo dopravy čr [online]. http://www.mdcr.cz. [15] eu. focus project [online], 2012. http://www.focusproject.eu/documents/14976/ -5d763378-1198-4dc9-86ffc46959712f8a. [16] fema. guide for all-hazard emergency operations planning, 1996. [17] us. us critical infrastructure conception, 2001. 58 www.ema.gov.au www.usfa.fema.gov http://www.mdcr.cz http://www.focusproject.eu /documents /14976/-5d763378-1198-4dc9-86ffc46959712f8a http://www.focusproject.eu /documents /14976/-5d763378-1198-4dc9-86ffc46959712f8a acta polytechnica ctu proceedings 5:51–58, 2016 1 introduction 2 current knowledge on transportation infrastructure 3 transportation system of the czech republic 4 method for determining the transportation infrastructure criticality 5 results – the criticality rates for transport infrastructure 6 proposals for transport infrastructure safety improve 7 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2018.20.0065 acta polytechnica ctu proceedings 20:65–72, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app isentropic efficiency of centrifugal compressor working with real gas jiří oldřich howden čkd compressors s.r.o., klečákova 347/5, 190 00, praha 9, czech republic correspondence: jiri.oldrich@howden.com, oldrich.jiri@seznam.cz abstract. the contribution deals with calculation of isentropic efficiency and also with calculation of isentropic process of real gas or gaseous mixtures. the method is based on numerical solution of basic definitional equation of isentropic process and equation of isentropic efficiency with direct implementation of real gas equation of state (eos). keywords: efficiency, isentropic, real gas, compressor. 1. introduction isentropic efficiency is an important parameter widely used in thermodynamic design of centrifugal compressor. reversible adiabatic process identical to isentropic process is used as reference process. the isentropic process is a process to which great attention is devoted in literature. in some articles certain inaccuracies are used repeatedly. for example dependence of heat capacities on temperature is ignored. very often poisson equations are used for calculation of real gases. isentropic exponents that are generally dependent on temperature and pressure are often substituted by mean values. adiabatic and isentropic processes are considered synonymous. these inaccuracies can generate noticeable errors in calculations. the aim of this contribution is to clarify the above mentioned inaccuracies and to suggest thermodynamically correct procedures that are appropriate for pc’s. in the first part of contribution equation for some simple special cases is described very shortly. the second more extensive part deals with techniques for various combinations of input and output variables. all these techniques are based on behavior of real gases and are independent on real gases eos used. 2. isentropic efficiency efficiency of compressor that does not exchange heat with the environment (compressor without intercooling or uncooled compressor section) is calculated by comparing the work that was actually spent on compression and the appropriate reference process. reference process is usually, depending on the compression process in the compressor, an isentropic process [1– 4], a polytropic process [2, 3, 5, 6], or a reversible isothermal process [2, 3, 7]. we will only deal with a reversible isentropic process and isentropic compression in the following text. in the actual process that occurs in the compressor, losses occur in the flowing gas and, as a result of it, some part of the kinetic energy is converted into heat. this process leads to an increase in temperature and figure 1. isentropic (area acdefa) and actual work (area gbcdefg) in t-s diagram when compressed from pressure p1 to pressure p2. entropy of compressed gas. a small part of the generated heat is discharged from the compressor casing to the surroundings. the process that is going on in the actual compressor is irreversible. the final temperature of the gas after its compression is always higher than temperature of gas after isentropic compression. in the diagram h-s in figure 1 the reversible isentropic compression work is represented by the area acdefa and the actual work spent on compression from pressure p1 to the pressure p2 = p2a = p2s is represented by the area gbcdefg. the work actually spent is irreversible and can be divided into isentropic work, compression losses (area abgfa) and increase of compression work due to losses (area abca) ∆ha = h2a −h1. (1) isentropic efficiency is defined as a relation between reversible isentropic work and actual work ηs = ∆hs ∆ha = h2s −h1 h2a −h1 . (2) short repetition of basic principles [2, 3, 8]. • adiabatic (isentropic) efficiency does not depend only on amount of losses. it depends on amount 65 http://dx.doi.org/10.14311/app.2018.20.0065 http://ojs.cvut.cz/ojs/index.php/app jiří oldřich acta polytechnica ctu proceedings figure 2. isentropic and actual processes in h −s diagram. of additional increase in compression work due to losses expressed by the area acba, too. • there are losses in gas flow in the actual process that occurs in compressor, and as a result, a part of the kinetic energy is converted into thermal energy and hence an increase in temperature and entropy of compressed gas. • at the same quality of compression mean value of additional increasing in compression work increases with increasing of total pressure. • at the same quality of compression the adiabatic efficiency will be at higher pressure ratio lower than at lower pressure ratio. • temperature of gas after actual compression is always higher than temperature after isentropic compression (see figure 2). 3. equation of isentrope adiabatic process is defined by the equation dq = 0, (3) where q is heat input into system ds = dq t = 0. (4) reversible adiabat is identical to isentrope. in irreversible process is ds > dq t = 0. (5) adiabat and isentrope are not identical. we rewrite equation (4) into form ds = ( ∂s ∂t ) v dt + ( ∂s ∂v ) t dv = = cv t dt + ( ∂p ∂t ) v dv = 0, (6) where p is the pressure of system and cv is heat capacity, which is derivative of internal energy with respect to temperature at constant volume cv = ( ∂u ∂t ) v . (7) that generally depends on temperature and on volume. for ideal gas the heat capacity cv = c0v depends only on temperature. equation of isentrope is obtained by integrating the relationship (6). its specific form depends on the chosen equation of state from which we determine (∂p/∂t)v and the chosen equation for the temperature dependence of c0v . 4. reversible process, ideal gas, constant c0v (perfect gas) this simplest case leads to the known poisson equations. equation (6) we rewrite into form c0v t dt = − r v dv, (8) where derivative (∂p/∂t )v is found from equation of state written for one mole pv = rt. (9) after integration of eq. (8) and after modification we receive well known poisson equation in variables t −v tv k−1 = const., (10) where k = (c0v + r)/c 0 v . poisson equations in variables t−p and p−v we obtain by inputting of volume v or temperature t from equation of state (9) tp 1−k k = const., (11) pv k = const. (12) 5. reversible process, general equation of isentrope from eq. (4) we obtain general equation of isentrope s(t,v ) = const. (13) then the problem is reduced to standard calculation of entropy for known (chosen) equation of state and equation of temperature dependence of c0v (figure 3). in the following text we will use the molar density ρ = 1/v instead of the volume as an independent variable. we will write the equation of the isentrope in the form s(t2,ρ2) −s(t1,ρ1) = 0, (14) where index 1 denotes inlet state and index 2 denotes end state after isentropic compression (figure 2). 66 vol. 20/2018 isentropic efficiency of centrifugal compressor working with real gas figure 3. dependence of c0p on temperature for several common used gases [9]. 6. methods of calculation of isentropic process and isentropic efficiency before we come to the description of the solution of each type of calculation, it will be useful to mention several methods of solving the isentropic process. • in the simplest case, we can assume that we are compressing the ideal gas. the calculation will be very easy and very inaccurate. • entropy was often read from diagrams or tables in the past. the accuracy was limited by the size of the diagram. however, this method is completely inappropriate for calculating gaseous mixtures and cannot be used on pc. • another of the approximate methods of calculating the isentropic process is the method based on the relations valid for the ideal gas, i.e. on the poisson equations (10), (11) and (12). isentropic exponent is calculated using of real gas eos and general relationships. kp,v = − v p cp cv ( ∂p ∂v ) t , (15) kp,t = 1 1 − p cp ( ∂v ∂t ) p , (16) kt,v = 1 + v cv ( ∂p ∂t ) v . (17) usually we do not know all parameters of end point of isentropic change (figure 2). thus we have to use value of isentropic exponent calculated for mean pressure and temperature. we estimate parameters in the point of interest from the behavior of the ideal gas. this method is suitable for pc and is appropriate for calculation of gas mixtures. even in this case, we do not have enough accurate results. as an illustration of the inappropriate use of the relationships derived for the ideal gas in combination with the isentropic exponent calculated by means of the real gas equation of state from the relationship (15), the calculation of the isentropic change from the relation can serve calculation of isentropic enthalpy from equation ∆hs = p1v1 kp,v kp,v − 1  (p2 p1 )kp,v −1 kp,v − 1   . (18) there are a number of other methods described in literature but further we focus only on methods which use the solution of a system of nonlinear equations, one of which is always a chosen real gas equation of state and one the requirement for zero change of entropy in transition from state 1 to 2 (figure 2). the following paragraphs provide algorithms of solution for individual types of calculations. in their design, particular emphasis was placed on working reliably and quickly. for every combination of input parameters the individual algorithm for solution of system of nonlinear equations was designed. important parts of every algorithm are relations for estimation of the initial approximations. in all cases, the calculation can be divided into a part independent of the real gas eos and the equation for the c0p = c0p(t ) of the ideal gas and a part dependent on these equations. the suitable method for solution of this system is newton’s method. functions fi that are used in the next detailed description of algorithms are shown in table 3. to construct the jacobi matrix of functions fi that we need for solution by the newton method, we need to know the partial derivatives of fi. these derivatives are shown in table 4. the abovementioned derivatives can also be obtained by numerical derivation of functions fi. in table 5 there are relations for calculation of entropy and enthalpy of real gas and relations for calculation of dimensionless quantities q [10]. any real gas eos can be chosen for the following calculations. dimensionless quantities q for chosen eos can be found in literature or can be derived from relations described in table 5. 7. calculations based on the real gas eos in practical calculations, we need to find unknown parameters when we know the initial state and two parameters of discharge state. the combination of input and output parameters is shown in table 1. in case that isentropic efficiency has value one i.e. ηs = 1 it is true p2s = p2a, t2s = t2a and ∆hs = ∆ha, then items 1, 2 and 3 from table 1 show combination of input and output parameters of isentropic change of thermodynamic state see table 2. 67 jiří oldřich acta polytechnica ctu proceedings input output 1 p1, t1, p2s t2s, ∆hs 2 p1, t1, t2s p2s, ∆hs 3 p1, t1, ∆hs p2s, t2s table 1. combination of output and input parameters for isentropic process. input output 1 p1, t1, p2a, ηs t2a, ∆ha 2 p1, t1, p2a, ∆ha t2a, ηs 3 p1, t1, t2a, ηs p2a, ∆ha 4 p1, t1, t2a, ∆ha p2a, ηs 5 p1, t1, p2a, t2a ∆ha, ηs table 2. combination of output and input parameters. 7.1. example of the solution by newton’s method all combinations of parameters are mentioned in [4]. for one of them we will show the solution. we will compute isentropic temperature t2 = t2s and isentropic change in enthalpy ∆h = ∆hs from known parameters p1, t1 and p2 (figure 2). it is variant 1 from table 1. in this case we assume that the isentropic efficiency is equal to one ηs = 1. since this is the isentropic process, the original system of four equations is now simplified to only two equations f1 = 0, f2 = 0. when we use general equation of isentropic process s1 − s2 = 0 (figure 2) and general expression of real gas eos we obtain the system of two nonlinear equations for two variables t2 and ρ2 f1 = s1(t1,ρ1) −s2s(t2s,ρ2s) = 0, (19) f2 = p2s ρ2srt2s −z2s(t2s,ρ2s) = 0. (20) this system of equations will be solved by newton’s method. because it is known and very often used method, we will show only an matrix notation of linear equations for increments in ∆t and ∆ρ[ ∂f1 ∂t2s ∂f1 ∂ρ2s ∂f2 ∂t2s ∂f2 ∂ρ2s ] · [ ∆t ∆ρ ] = [ −f1 −f2 ] (21) desired parameters t2 and ρ2 can be obtained from relationships t (k+1) 2s = t (k) 2s + ∆t, (22) ρ (k+1) 2s = ρ (k) 2s + ∆ρ. (23) the calculation is completed by fulfilling of the requiremen max(|δt|; |δρ|) ≤ � where δt and δρ are relative deviations and � is chosen value. detail description of solution for this example: for inlet condition p1, t1 we determine values s1 and h1 from real gas eos next we calculate c0v c0p = c 0 p(t1, −→x ), c0v = c 0 p −r. (24) next step is solution of system of equations f1 = 0, f2 = 0 by using of newton’s method. initial approximations are relations valid for ideal gas t (0) 2s = t1 ( p2s p1 )c0p −c0v c0 p , (25) ρ (0) 2s = p2s rt (0) 2s . (26) as the results we obtain t2s and ρ2s. from general form of eos we obtain compressibility factor in point 2s (figure 2) z2s = p2s rt2sρ2s (27) and enthalpy h2s = h2s(t2s,ρ2s,−→x ). (28) finally we calculate change in enthalpy ∆h = h2 −h1, (29) where h2 = h2s. 7.2. calculation of t2s, ∆ha from known parameters p1, t1, p2a and ηs (figure 2, table 2 item 1) we solve the system of equations f1 = 0, f2 = 0, f4 = 0 and f5 = 0 for unknown parameters t2s, ρ2s, t2a, ρ2a. initial approximations are t (0) 2s = t1 ( p2a p1 )c0p −c0v c0 p , (30) ρ (0) 2s = p2a rt (0) 2s , (31) t (0) 2a = t (0) 2s −t1 ηs + t1, (32) ρ (0) 2a = p2a rt (0) 2a , (33) where c0p is c0p = c 0 p(t1, −→x ), c0v = c 0 p −r. (34) change in enthalpy ∆ha we obtain from equation ∆ha = h2a −h1, (35) where h2a = f(t2a,ρ2a,−→x ). (36) 68 vol. 20/2018 isentropic efficiency of centrifugal compressor working with real gas 7.3. calculation of t2a, ηs from known parameters p1, t1, p2a and ∆ha (figure 2, table 2 item 2) we solve the system of four equations f1 = 0, f2 = 0, f5 = 0, f6 = 0 for unknown parameters t2s, ρ2s, t2a, ρ2a. initial approximation t (0)2s we obtain from eq. (30), ρ (0) 2s from eq. (31) t (0) 2s = t1 + ∆ha c0p (37) and ρ(0)2a from eq. (33). isentropic efficiency ηs we calculate from eq. (2), where enthalpy h1 we know from calculation in initial point 1 and h2a = f(t2a,ρ2a,−→x ). 7.4. calculation of p2a, ∆ha from known parameters p1, t1, t2a and ηs (figure 2, table 2 item 3) calculation is based on solving of three equations f1 = 0, f4 = 0 and f8 = 0 for unknown parameters t2s, ρ2s and ρ2a. the first approximations are t (0) 2s = ηs(t2a −t1) + t1, (38) ρ (0) 2s = ρ1 ( t2s t1 ) c0v c0p −c 0 v , (39) ρ (0) 2a = t (0) 2s ρ (0) 2s t2a , (40) where c0p see eq.(34). after every step of the newton’s method the k-th approximation of the compressibility factor z(k)2a is found from real gas equation of state and then the k-th approximation of the pressure p(k)2a is calculated from equation p (k) 2a = z (k) 2a rt2aρ (k) 2a . (41) from equality p2s = p2a (see figure 2), we can calculate z (k) 2s = z (k) 2a t2aρ (k) 2a t (k) 2s ρ (k) 2s . (42) change in enthalpy ∆ha is determined from eq. (1), where h2a = f(t2a,ρ2a,−→x ). 7.5. calculation of p2a, ηs from known parameters p1, t1, t2a and ∆ha (figure 2, table 2 item 4) to calculate unknown parameters t2s, ρ2s and ρ2a system of three equations f1 = 0, f6 = 0 and f8 = 0 is solved. for the first approximation of temperature we choose t (0)2s = t2a. density ρ (0) 2s we obtain from equation (40), ρ(0)2a = ρ (0) 2s and z (0) 2a = z (0) 2s = 1. after every figure 4. ideal gas temperature t2s calculated from eq. (30) – blue curve and real gas temperature t2s – red curve. inlet conditions propane p1 = 550000 pa, t1 = 293.15 k. figure 5. ideal gas isentropic work ∆hs calculated from eq. (43) – blue curve and real gas isentropic work ∆hs – red curve. inlet conditions propane p1 = 550000 pa, t1 = 293.15 k. step of newton’s method compressibility factors z(k)2a and z(k)2s are calculated similar way like in previous case i.e. by solution of real gas eos and by using of equations (41) a (42). finally isentropic efficiency ηs we obtain from eq. (2), where h1 is known from calculation of thermodynamic properties in point 1 (figure 2) and h2s = f(t2s,ρ2s,−→x ). 7.6. calculation of ∆ha, ηs from known parameters p1, t1, p2a and t2a (figure 2, table 2 item 5) to calculate unknown parameters t2s, ρ2s and ρ2a system of three equations f1 = 0, f2 = 0 and f5 = 0 is solved. first approximation of temperature t (0)2s we obtain from eq. (30), ρ(0)2s from eq (31) and ρ (0) 2a from eq. (33). change in enthalpy ∆ha we calculate by using of eq. (1), where h2a = f(t2a,ρ2a,−→x ). isentropic efficiency ηs we finally obtain from eq. (2), where h2s = f(t2s,ρ2s,−→x ). 69 jiří oldřich acta polytechnica ctu proceedings figure 6. relative deviation of isentropic work calculated for real and ideal gas. 8. example as an example the calculation of temperature and isentropic work after isentropic change was chosen. in this example compression of propane from state 1 to state 2s (see figure 2) is calculate by two methods. in the first case above described method based on the real gas is used and in the second case method based on an ideal gas. parameters in point 1 are pressure p1 = 550000 pa, temperature t1 = 293.15 k. the third parameter is gas pressure p2s which varies in the interval 800000 to 2500000 pa. for calculation of real gas isentropic change the method described in paragraph 7.1 was used. this method is equivalent to method of paragraph 7.2 when ηs = 1. for calculation of temperature t2s of ideal gas the equation (30) was for p2a = p2s = p2. isentropic change of enthalpy was calculated from eq. (18) in form ∆hs = z1rt1 k k − 1 [( p2 p1 )k−1 k − 1 ] , (43) where z1 and k were calculated from bwr eos. results of all calculations are on figures 4, 5 and 6. 9. conclusion the methods described in this contribution exactly solved problem of calculation of the isentropic efficiency for various input parameters. in case that efficiency is equal zero above described procedures are suitable for calculation of isentropic process. the above described methods advantageously use the dimensionless variables defined in table 5 and real gas eos. it is evident from the work that some approximate procedures can lead to inaccurate results. methods described here together with method described in [6] and in [7] and suitable real gas eos form the basic tools for thermodynamic calculations of compressors or others flow machines. list of symbols t thermodynamic temperature [k] p pressure [pa] r molar gas constant [j/(mol k)] z compressibility factor [–] ρ molar density [mol/m3] v molar volume [m3/mol] k isentropic exponent [–] cv molar heat capacity at constant volume [j/(mol k)] cp molar heat capacity at constant pressure [j/(mol k)] h molar enthalpy [j/(mol)] s molar entropy [j/(mol k)] −→x vector of molar fractions of gas mixture [–] η efficiency [1] t0 standard temperature [k] p0 standard pressure [pa] eos equation of state indexes: s isentropic a actual 1 inlet, input 2 discharge, output id ideal gas 0 thermodynamic quantities of ideal gas (upper index) references [1] d. a. kouremenos, x. k. kakatsios. the three isentropic exponents of dry steam. forschung auf dem gebiete des ingenieurwesens 51:117–122, 1985. doi:10.1007/bf02558416. [2] d. misárek. turbokompresory (turbocompressors). sntl, 1963. in czech. [3] k. h. lüdtke. process centrifugal compressors. springer, berlin, heidelberg, 2004. doi:10.1007/978-3-662-09449-5_9. [4] j. oldřich, j. novák, a. malijevský. adiabatický děj v technické praxi, (adiabatic pro-cess in technical practice). strojírenství 40(2):69–76, 1990. in czech. [5] j. m. schultz. the polytropic analysis of centrifugal compressors. journal of engineering for power 84:69, 1962. doi:10.1115/1.3673381. [6] j. oldřich. advanced polytropic calculation method of centrifugal compressor. 2010. doi:10.1115/imece2010-40931. [7] j. oldřich. calculation of the isothermal efficiency of a cen-trifugal compressor compressing the real gas. in proceedings of conference “turbostroje 2016”. plzeň, czech republic, 2016. [8] b. eckert. axialkompressoren und radialkompressoren. springer-verlag, berlin, 1953. doi:10.1007/978-3-642-52720-3. [9] p. j. linstrom, w. g. mallard. nist chemistry webbook, nist standard refer-ence database number 69. doi:10.18434/t4d303. [10] j. p. novák, a. malijevský, j. pick. použití bezrozměrných veličin při výpočtu termo-dynamických funkcí, (use of dimensionless quan-tities for calculating of thermodynamic functions from equations of state). chemické listy 73:1178–82, 1979. in czech. 70 http://dx.doi.org/10.1007/bf02558416 http://dx.doi.org/10.1007/978-3-662-09449-5_9 http://dx.doi.org/10.1115/1.3673381 http://dx.doi.org/10.1115/imece2010-40931 http://dx.doi.org/10.1007/978-3-642-52720-3 http://dx.doi.org/10.18434/t4d303 vol. 20/2018 isentropic efficiency of centrifugal compressor working with real gas f1 = s1(t1,ρ1) −s2(t2s,ρ2s) f2 = p2s ρ2srt2s −z2(t2s,ρ2s) f3 = ∆hs − [h2(t2s,ρ2s) −h1(t1,ρ1)] f4 = ηs[h2(t2a,ρ2a) −h1(t1,ρ1)] − [h2(t2s,ρ2s)− −h1(t1,ρ1)] f5 = p2s ρ2art2a −z2(t2a,ρ2a) f6 = [h2(t2a,ρ2a) −h1(t1,ρ1)] − ∆ha f7 = ηs∆ha − [h2(t2s,ρ2s) −h1(t1,ρ1)] f8 = z2(t2a,ρ2a)t2aρ2a −z2(t2s,ρ2s)t2sρ2s table 3. functions fi. ∂f1 ∂t2s = r t2s (1 + 2qu2s + qc2s) − ∑ xic 0 pi(t2s) t2s ∂f1 ∂ρ2s = rqt2s ρ2s ∂f2 ∂t2s = − qt2s t2s ∂f2 ∂ρ2s = − qρ2s ρ2s ∂f2 ∂t2a = ∂f2 ∂ρ2a = ∂f6 ∂t2s = ∂f6 ∂ρ2s = ∂f7 ∂ρ2a = 0 ∂f3 ∂t2s = ∂f4 ∂t2s = ∂f7 ∂t2s = −r(qt2s − 2qu2s −qc2s− −1) − ∑ xic 0 pi(t2s) ∂f3 ∂ρ2s = ∂f4 ∂ρ2s = ∂f7 ∂ρ2s = rt2s(qt2s −qρ2s) ρ2s ∂f4 ∂t2a = ηs[r(qt2a − 2qu2a −qc2a − 1)+ + ∑ xic 0 pi(t2s)] ∂f4 ∂ρ2a = ηs rt2a(qρ2a −qt2a) ρ2a ∂f1 ∂t2a = ∂f1 ∂ρ2a = ∂f5 ∂t2s = ∂f5 ∂ρ2s = 0 ∂f5 ∂t2a = − qt2a t2a ∂f5 ∂ρ2a = − qρ2a ρ2a ∂f6 ∂t2a = r(qt2a − 2qu2a −qc2a − 1)+ + ∑ xic 0 pi(t2s) ∂f6 ∂ρ2a = rt2a(qρ2a −qt2a) ρ2a ∂f8 ∂t2s = −ρ2sqt2s ∂f8 ∂ρ2s = −t2sqρ2s ∂f8 ∂t2a = ρ2aqt2a ∂f8 ∂ρ2a = t2aqρ2a table 4. derivative of functions fi. 71 jiří oldřich acta polytechnica ctu proceedings z = z(t,ρ) = p rtρ s(t,ρ,−→x ) = ∑ x1 [ s0i (t0) + ∫ t t0 c0pi(t) t dt ] − −r ln zrtρ p0 −r ∑ xi ln xi + r(ln z −qf −qu ) h(t,ρ,−→x ) = ∑ x1 [ h0i (t0) + ∫ t t0 c0pi(t)dt ] + +rt(z − 1 −qu ) qρ = z + ρ ( ∂z ∂ρ ) t qf = ∫ ρ 0 (z − 1)d(ln ρ) qt = z + t ( ∂z ∂t ) ρ qu = t ( ∂qf ∂t ) ρ qc = t 2 ( ∂2qf ∂t 2 ) ρ table 5. relations for calculation of entropy and enthalpy of real gas and relations for calculation of dimensionless quantities q [10]. 72 acta polytechnica ctu proceedings 20:65–72, 2018 1 introduction 2 isentropic efficiency 3 equation of isentrope 4 reversible process, ideal gas, constant cv0 (perfect gas) 5 reversible process, general equation of isentrope 6 methods of calculation of isentropic process and isentropic efficiency 7 calculations based on the real gas eos 7.1 example of the solution by newton's method 7.2 calculation of t2s, ha from known parameters p1, t1, p2a and s (figure 2, table 2 item 1) 7.3 calculation of t2a, s from known parameters p1, t1, p2a and ha (figure 2, table 2 item 2) 7.4 calculation of p2a, ha from known parameters p1, t1, t2a and s (figure 2, table 2 item 3) 7.5 calculation of p2a, s from known parameters p1, t1, t2a and ha (figure 2, table 2 item 4) 7.6 calculation of ha, s from known parameters p1, t1, p2a and t2a (figure 2, table 2 item 5) 8 example 9 conclusion list of symbols references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0039 acta polytechnica ctu proceedings 13:39–42, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app on development of the interactive design tool based on isogeometric analysis edita dvořáková∗, bořek patzák czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: edita.dvorakova@fsv.cvut.cz abstract. isogeometric analysis is a new concept of finite element method which has been proposed to bridge the gap between the cad systems and the fem solvers. in isogeometric analysis, the same basis functions, typically splines or nurbs, are used for geometry description as well as for approximation of unknowns and thus the same model can be shared between cad and iga systems. this results in a higher accuracy and overall efficiency of the analysis. many isogeometric elements have been already proposed and implemented into the existing finite element solvers, but the automatic connection with a cad system is usually missing. our goal is to develop such a connection and to provide a tool which would interactively run the analysis when the model changes. this approach can enormously enhance the design process as it can provide the basic knowledge about the structural behavior already in conceptual design phase. keywords: curved beam element, discrete shear gap, interactive design, isogeometric analysis. 1. introduction finite element solvers are the integral part of a nowadays structural design. once the architectural model is developed in cad (computer-aided design) system, the structure is further analyzed in fem (finite element method) solver. there are two independent models (cad model and fem model) and thus when one of them changes during the design process the other has to be correspondingly (and usually manually) adjusted, leading to the higher time-consumption. moreover, the inaccuracies in geometry description are usually introduced as the cad and fem models use different functions for geometry representation. the standard finite element method usually uses polynomial basis functions, while the cad systems are based on splines. the use of spline functions for geometry and unknown approximations in finite element analysis enables to share one model between architectural design and structural analysis. this idea has been proposed by huges et al. [1] and it is called isogeometric analysis (iga). since 2005, when isogeometric analysis has been introduced, many researchers have been focusing on its development and the results proved the advantages of iga over standard fem in many fields. our focus is placed on its application for the analysis of curved beams as the curved geometries can especially benefit from the isogeometric approach. we use the formulation of curved beam element presented by bouclier et al. [2] with the locking treatment based on discrete shear gap (dsg) method [3, 4]. the element has been implemented into oofem finite element solver [5] and our main goal is to develop a connection between cad system and the solver to provide a tool encapsulating the analysis within the architectural design environment. 2. curved beam element the presented two-dimensional curved beam element is based on timoshenko beam theory. it has three independent unknowns, axial displacement ut, transverse displacement un and rotation θ. a strain-displacement matrix b defined as ε = br, (1) where ε = {εm,γs,χb}t and r = {ut,un,θ}t , is derived from the relations for membrane strain εm, transverse shear strain γs, and bending strain χb εm = u′t − un r , γs = ut r + u′n −θ, χb = θ′, (2) where s ∈ 〈0,l〉 runs along the midline of a beam and prime indicates a derivative with respect to s. by omitting the terms which are divided by radius of curvature r the straight beam formulation is obtained. a material matrix d results from n = eaεm, q = kgaγs, m = eiχb, (3) where n,q and m are axial force, transverse shear force and bending moment, respectively. young’s modulus e, shear modulus g, shear coefficient k, area a and moment of inertia i are the material and 39 http://dx.doi.org/10.14311/app.2017.13.0039 http://ojs.cvut.cz/ojs/index.php/app edita dvořáková, bořek patzák acta polytechnica ctu proceedings cross-section characteristics. a stiffness matrix is then calculated as k = ∫ l 0 bt db ds. (4) to evaluate the stiffness matrix the gaussian quadrature involving p + 1 gauss points over each knotspan is used, where p is a degree of approximation. 2.1. nurbs-based formulation in case of iga, the langrange polynomial basis functions used in fem are replaced by splines. the presented element is formulated using non-uniform rational b-splines (nurbs) [6], which is the most widespread and developed technology in cad industry. a nurbs curve is defined by its degree, control points with weights, and knotvector. nurbs basis functions are generated by weighting the b-splines functions, which are piecewise polynomial functions and are a special subset of nurbs functions. nurbs curve is given as the linear combination of basis functions and control points coordinates. the knotvector is a set of non-decreasing parametric coordinates (knots) which influence the mapping. the knots divide a parametric space into knotspans (usually referred to as “elements” in iga) and each knotspan is influenced only by limited number of control points. therefore the knotspan defines which control points affect particular part of the domain. see figure 1 for an example of a nurbs curve geometry. very important aspect of isogeometric analysis is a non-interpolatory nature of the basis functions, as the nurbs do not satisfy kronecker-delta property. another distinct feature of the nurbs basis functions is the higher inter-element continuity. while only c0 continuity is provided by traditional polynomial functions, in case of iga up to cp−1 continuity can be achieved. by repeating knots in a knotvector, the continuity at that knot can be artificially reduced. 2.2. locking treatment shear locking is a well-known problem of standard timoshenko beam elements and this phenomena persists also in isogeometric formulation. this problem occurs because the displacements and rotation are treated independently and approximated by the functions of the same order. from eq. 2 for shear and bending strains it can be seen that formula for bending strain results in lower order term than formula for shear strain, but actually this should be vice-versa. also the field inconsistency within the term for shear strain causes that zero shear strain cannot be satisfied along entire patch when the same order interpolation of unknowns is used. several locking removal techniques have been proposed to unlock isogeometric beam element including reduced integration, b̄-method, dsg method, see [2, 3] for the references. in this study, the dsg-based formulation is used. while this method can increase computational cost (as well as b̄-method), it has a big advantage over reduced integration: for the reduced integration the recovered strains can be evaluated only at gauss points, dsg method enables to evaluate them along the entire patch. discrete shear gap method has been originally developed for the standard finite elements [3] but echter and bischoff [4] have extended its use also for isogeometric analysis. the approach can be divided into several steps yielding the modified b matrix. the idea is not to satisfy the equation γs = ut r + u′n −θ (5) pointwise, but in integral sense. the shear contribution uγ hi n (called “shear gaps”) to the deflection is obtained by integration of eq. 5 as uγ hi n = ∫ si 0 γhs ds = ∫ si 0 ut r + u′n −θ ds = b dsgr, (6) at collocation points si calculated as greville abscissa of the control points [6]. modified transverse displacements are interpolated using nurbs basis functions ni uγ modh n = n∑ i=1 niũ γhi n . (7) please note, that discrete shear gaps ũγ hi n are noninterpolatory in isogeometric analysis. to expressed the discrete shear gaps by values at the control points the transformation matrix a is derived{ uγ h n } = a { ũγ h n } , aij = nj(si), (8) where { uγ h n } are interpolatory values of shear gaps at control points and nj(si) is the jth-basis function evaluated at ith-collocation point. the modified shear strain is then given as γmod h s = n∑ i=1 n′iũ γhi n . (9) combining eqs. 6-9 results in the modified straindisplacement matrix b b = n′a−1bdsg, (10) which is used to evaluate stiffness matrix k (eq. 4). 3. interactive design tool thanks to the isogeometric formulation, the presented element can be easily connected to cad system. the possibility of seamless connection between cad and fea is the major benefit of iga. this goal can be achieved even without use of isogeometric approach (see e.g. [7]), nevertheless the costly transformations 40 vol. 13/2017 on development of the iga based interactive design tool knots control points s = l s = 0 s ξ ξ1 = 0 ξ2 = 0.5 ξ3 = 1 knot span patch y x figure 1. description of nurbs geometry. between cad and fem are necessary in such a workflow. firstly, the geometrical model is completed by analysis data, such as loads, boundary conditions and material characteristics, then the model is transformed into suitable computational model and the analysis mesh is generated, and finally the analysis is performed in fem solver. to visualize the results within the cad system the model has to be transformed back to the original model. with the use of isogeometric analysis, once the model is completed with analysis data, the analysis can be performed directly [8] and no transformation is needed for both the analysis and the visualization. for our purposes, the use of cad system rhinoceros (rhino) [9] has been chosen. the geometry representation of rhino is based on nurbs, moreover rhino enables plug-in development and use of built-in tool grasshopper [10]. grasshopper is a visual programming interface within rhino which can directly access nurbs geometry. in addition, visual programming is intelligible even to a user with no programming background, so the user can define additional analysis data at this level. the workflow of the tool is illustrated in figure 2. firstly, the nurbs geometry specified by a user in rhino is passed to the grasshopper, where the model is completed with boundary conditions and material and crosssection characteristics. currently, these data are provided in the text format similar to one used in oofem, but in the next phase of the development, the special plug-ins can be developed. once all the analysis data are collected, the python script within grasshopper together with other available grasshopper tools directly passes these data to oofem and run the analysis. when the analysis is finished, the output file is uploaded back to the grasshoper, where the visualization of the results is carried out. schematic illustration of grasshopper environment with results visualized in rhino can be seen in figure 3. very important feature of the developed tool is, that the grasshopper (and in turn also oofem) recomputes the results immediately when something changes. no matter whether the user adjusts the geometry model results analysis data geometry material, crosssection boundary conditions analysis results visualisation rhinoceros 3d grasshopper grasshopper oofem figure 2. interactive design tool. geometry or changes the boundary conditions, the results are interactively updated. this provides the immediate knowledge about behaviour of the designed structure and eases the process of finding optimal design. in the future work the designer will receive graphical indication of sections where load-bearing capacity or stability limits are achieved. such an approach would allow to establish viable conceptual design from both architectural and structural point of view without the necessity of understanding details of structure behaviour. 4. conclusions the formulation of the curved beam element with locking treatment has been presented and the element has been implemented into oofem finite element code. thanks to the isogeometric formulation, the automatic connection of cad system and the analysis has been provided or modified. the developed tool enables to display the results in rhino immediately 41 edita dvořáková, bořek patzák acta polytechnica ctu proceedings figure 3. schematic illustration of grasshopper environment with results visualized in rhino. the example corresponds to a cantilever circular arc beam (black) subjected to horizontal and vertical load at its tip. the deformed geometry is in red and the bending moment is in blue. after the geometry and analysis data are provided. this gives the user instant knowledge about behaviour of the designed structure and enables to see differences of alternative designs. acknowledgements the financial support of this research by the grant agency of the czech technical university in prague (sgs project no. sgs17/043/ohk1/1t/11) is gratefully acknowledged. references [1] t. j. r. hughes, j. a. cottrell, y. bazilevs. isogeometric analysis: cad, finite elements, nurbs, exact geometry and mesh refinement. comput methods appl mech engrg 194:4135–4195, 2005. [2] r. bouclier, t. elguedj. locking free isogeometric formulations of curved thick beams. comput methods appl mech engrg 245-246:144–162, 2012. [3] k. bletzinger, m. bischoff, e. ramm. a unified approach for shear-locking-free triangular and rectangular shell finite elements. computers and structures 75:321–334, 2000. [4] r. echter, m. bischoff. numerical effiency, locking and unlocking of nurbs finite elements. comput methods appl mech engrg 199:374–382, 2010. [5] b. patzák. oofem project home page, 2017. http://www.oofem.org. [6] l. piegl, w. tiller. the nurbs book. springer-verlag berlin heidelberg, new york, 1997. [7] l. svoboda, j. novák, l. kurilla, j. zeman. a framework for integrated design of algorithmic architectural forms. advances in engineering software 72:109–118, 2014. [8] m. hsu, c. wang, a. herrema, et al. an interactive geometry modeling and parametric desigh platform for isogeometric analysis. computers and mathematics with applications 70:1481–1500, 2015. [9] rhinoceros, 2017. http://www.rhino3d.com/. [10] grasshopper, 2017. http://www.grasshopper3d.com/. 42 http://www.oofem.org http://www.rhino3d.com/ http://www.grasshopper3d.com/ acta polytechnica ctu proceedings 13:40–43, 2017 1 introduction 2 curved beam element 2.1 nurbs-based formulation 2.2 locking treatment 3 interactive design tool 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.9.0008 acta polytechnica ctu proceedings 9:8–12, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app phase analysis of niti wire based on synchrotron 2d x-ray diffraction pattern in matlab martin dudra, ∗, daniel šimekb, jan drahokoupila, b a department of solid state engineerign faculty of nuclear sciences and physical engineering czech technical university in prague, trojanova 13 120 00 prague 2, czech republic b institute of physics , na slovance 2 182 21 prague, czech republic ∗ corresponding author: dudrmart@fjfi.cvut.cz abstract. this paper deals with simplified phase analysis of 2d diffraction patterns obtained using microfocused x-ray at various places of superelastic niti wire under combined load. for this purpose, matlab routines and functions were created, which carry out integration of the patterns in various sectors of debye ring, fitting of the patterns by a mixture of austenite and martensite (having neglected the r-phase) and successive evaluation of the quantitative phase content. the most crucial simplificating assumptions were: the neglection of the r-phase and the assumption of a) fibre texture or b) no texture. keywords: nitinol spring, combined load, x-ray diffraction, phase analysis, synchrotron. 1. introduction a superelastic nitinol wire under combined load was subjected to xrd measurements using microfocused beam at european synchrotron radiation facility (esrf) in grenoble. quantitative phase analysis at different points on the wire was expected to be carried out. as the number of outcoming images was large (12 000), there was need of their automated processing. moreover the sample was strongly textured and therefore the integration of debye rings was preferred to be done "per partes", that means separately in various sectors ("pieces of cake") in order to implement some texture corrections later. some matlab routines were created that fulfilled these specific demands. this paper briefly explains basic principles of the approach in use and presents results obtained from the routines. 2. material and methods 2.1. nitinol nitinol is a well-known shape memory alloy of a near equiatomic stoichiometric composition ni-ti. the shape memory in niti is based on so called martensitic transformation (mt) between cubic austenite (high-temperature phase) and monoclinic twinning martensite (low-temperature phase). mt can be induced by lowering the temperature below transition point or by applying sufficient mechanical stress onto austenite. there is a third phase called r-phase, which under certain conditions stands in between the former two. it is unstable and hardly preparable solely. [1] r-phase is trigonal. [2] 2.2. quantitative phase analysis quantitative phase analysis (qpa) responds to both questions "what?" and "how much?" when talking about a mixture of phases. the diffraction pattern outcoming from measurement on a polycrystalline sample in transmission geometry recorded on 2d detector looks like a set of concentric circles. each of the circles corresponds to one set of equivalent crystallographical planes of the sample. when the pattern is integrated in tangential direction 1d patterns are obtained for the integrated interval. position of the peak in 1d pattern (i.e. the radius of the corresponding circle) is important for qualitative phase analysis (i.e. "what?"), the integrated intensity of the peak determines the amount of phase in a mixture ("how much?"), hence it is used by quantitative phase analysis. there are many factors determining the real value of integrated intensity. let ihkl be an integrated intensity of crystallographical planes (hkl), then we can write [3] ihkl = k ·phkl ·lθ ·pθ ·aθ ·thkl ·ehkl ·|fhkl| 2 . (1) k is a scale factor, which is proportional to diffracting volume of the phase and therefore it is of great importance in qpa. other factors are multiplicity of planes phkl, lorentz factor lθ, polarization factor pθ, absorption factor aθ , texture correction thkl which will be discussed further on, extinction factor ehkl and important structure factor squared |fhkl| 2. the principle of qpa is very simple: to isolate the scale factors k for the phases in mixture from the equation 1. nevertheless, some of the other factors and corrections need not to be easily determinable. especially texture correction may be troublesome as texture (also called preferred orientation) description usually requires specific measurement and nontrivial data analysis. in this paper the authors tried to approximate the real texture by two simplificating cases separately: 8 http://dx.doi.org/10.14311/app.2017.9.0008 http://ojs.cvut.cz/ojs/index.php/app vol. 9/2017 phase analysis of niti wire based on synchrotron 2d xrd pattern mm 0 1 2 3 4 5 6 7 8 9 10 m m 0 0.5 1 figure 1. visualization of gridpoints showing diffraction on the sample. one pixel corresponds to one image, axes have equal scale. a) radially symmetric fibre texture and b) no texture at all. 2.3. experimental the sample was a coiled round wire of 0.2 mm diameter from superelastic alloy niti#1 [4]. this helical spring was elongated in uniaxial tension, which induced a complex stress state consisting of torsion and tension/compression due to bending of the spring. the tensile load was of 1.6 n. measurements were carried out at esrf, grenoble, at beamline 15a [5]. the wavelength of used radiation was λ = 0.0268128 nm. microfocused x-ray beam of rectangular cross-section 20 µm × 5 µm was used to scan in a rectangular grid of 12 000 points of which approximately 2 200 covered the sample. more precisely speaking it was a grid of 100 × 120 points and of measures 10 mm × 1.1 mm. so the step was 100 µm in the x direction and cca 9.1 µm in the y direction. in figure 1 the scanned area is to be seen. each pixel corresponds to one gridpoint (one measurement, one image). blue pixels represent measurements where only background was captured (low average intensity in the image) and the yellow pixels correspond to images capturing the sample diffraction pattern (average intensity was larger). the wire profile is clearly visible. 2.4. data processing before the data processing itself the wavelength and sample-detector distance were determined using two calibration images taken on calibrant ceo2 in different sample-detector distances. then the integration in polar angle was carried out. due to strong texture of the sample and as a macrodeformation is expected, the integration was carried out in nine sectors separately, as indicated in figure 2. as the structure of austenite and martensite is known [6] it is possible to estimate the positions of diffraction lines of these phases. there was designed such an algorithm, that tries to fit an analytical function pearson vii to place where we expect austenitic or martensitic line. the whole pattern is then superposition of these pearsons which are all fitted at once, hence it is a whole profile fitting approach, more specifically le bail fitting the latter means that the intensities of lines are fitted parameters (see parameter a1 below). it is possible to write pearson vii in figure 2. indication of nine sectors in which the integration was carried out. each sector has angular width of 20°. a following form [7] p vii(x) = a1 1 [1 + (x−a2)2a3] a4 . (2) parameters a1, a2, a3 and a4 define the height of the function, its position in x, width and shape respectively. in the work other form was used, but the above mentioned one is more clear. the parameter a2 was a function of lattice parameters and macrodeformation (voigt model of strain distribution was assumed). the parameter a3 was a function of crystallite size and microdeformation (accordingly to williamson-hall equation for width of a diffraction line). parameters a1 (for all lines) and a4 (for each phase austenite and martensite) were fitted directly. the area under pearson vii function is possible to express analytically from the parameters a1, a2, a3 and a4. this area represents the integral intensity of a given line. the fitting was performed using nonlinear least squares, by so called levenberg-marquardt method. the fitting was carried out in all the sectors at once. some parameters were related only to a specific sector (e.g. a height a1 of a particular line in a sector), others were fitted on basis of all sectors (e.g. 3 macrodeformation-describing parameters per phase). there were 60 linear parameters in each sector (height a1 of 50 lines of martensite and 7 lines of austenite plus 3 parameters describing background) and cca 10 other nonlinear parameters which was possible to exclude from fitting (these were related to a2, a3 and a4). altogether it makes cca 550 parameters to be fitted in one image. once the integral intensity of all lines (of austenite as well as martensite) is assumed to be known from fitted parameters, the left side of equation 1 is known and thus it is possible to start isolating scale factors k of both phases from the equation (if we know all other terms). in order to finish, the texture evaluation (in this case approximation) has to be carried out. the main difference between the fibre-texture approximation and no-texture approximation was especially in the weights, which were given to integral intensities 9 martin dudr, daniel šimek, jan drahokoupil acta polytechnica ctu proceedings mm 0 2 4 6 8 10 m m 0 0.2 0.4 0.6 0.8 1 aust mart mm 0 2 4 6 8 10 m m 0 0.2 0.4 0.6 0.8 1 aust mart figure 3. relative volume fractions of austenite and martensite with respect to the whole diffracting volume. in the left-hand figure the fibre-texture correction was applied, while on the right it was the no-texture approximation. the martensite is being represented by red channel, the austenite by the blue one. sin3/6 [angst-1] 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 in te n si ty [ ct s] #104 0.5 1 1.5 2 2.5 3 3.5 4 4.5 data fit figure 4. demonstration of fit in one sector. of lines from different sectors and in the number of lines which were taken into account. it is important to notice here that the r-phase was not taken into account in fitting. 3. results and discussion an example showing fit in one sector of image is in figure 4. each of the images taken on the sample (yellow pixels in figure 1) was processed individually and the ratios a/(a + m) and m/(a + m) were determined, where a resp. m stand for the volumes of austenite resp. martensite which contribute to diffraction, i.e. the volume fractions of the phases. the resulting volume fractions of both phases are in the figure 3. the strong prevalence of martensite is easily to be seen. four scans across the wire were chosen to demonstrate the results in detail, concretely scans 11, 36, 66 and 96 with no-texture correction. they have cca a quarter of spring coil period spacing in between, as mm 0 1 2 3 4 5 6 7 8 9 10 m m 0 0.2 0.4 0.6 0.8 1.1 11 96 6636 figure 5. indication of four scans which are presented in more detail. there were 100 scans in total, each of the indicated scans has its number denoted nearby. indicated in figure 5. each scan consists of cca 22 images. for the results themselves see figure 6. besides the diffracting volume of austenite and martensite there are two more informations in each graph. firstly a sum of austenite and martensite, which should follow an ellipse-like arch (the wire has more mass in the center than at the borders) and secondly illustrative semi-ellipse chosen so that its maximum value is slightly higher than maximum point of the sum of the phases. this is because the absorption correction was applied only partly during image integration, but it is still expected that the resulting intensity will be affected by absorption. the illustrative wire profile (semi-ellipse) serves only to easier comparison of the sum of volumes and the expectations. scans 36 and 96 mutually seem to be a mirror image of the other (with the mirror being placed vertically in the middle of each figure) in both we can see a decrease of sum of diffracting volumes close to the middle of the wire but shifted more to the right in scan 36 and to the left in scan 96. below this effect is explained by presence of r-phase. in figure 7 there is to be seen a central group of lines in the measured angle interval, where one line of austenite (110) and four lines of martensite are expected. obviously, there are two other lines which lay in between the austenitic line (110) and martensitic lines (1-11) and (002). the positions of austenitic and 10 vol. 9/2017 phase analysis of niti wire based on synchrotron 2d xrd pattern pattern no. 5 10 15 20 d iff ra ci n g v o lu m e [ a .u .] 0 0.05 0.1 0.15 scan no. 11 pattern no. 5 10 15 20 d iff ra ci n g v o lu m e [ a .u .] 0 0.05 0.1 0.15 scan no. 36 pattern no. 5 10 15 20 d iff ra ci n g v o lu m e [ a .u .] 0 0.05 0.1 0.15 scan no. 66 pattern no. 5 10 15 20 d iff ra ci n g v o lu m e [ a .u .] 0 0.05 0.1 0.15 scan no. 96 aust mart sum aust+mart wire profile figure 6. results of phase analysis for four scans (no-texture approximation). the y axis shows diffracting volumes, x shows the order number of the images in the given scan. blue color stands for the diffracting volume of austenite, the red one for the diffracting volume of martensite and the green one is a summation of the two. the black line indicates the theoretical wire mass profile. martensitic lines in the group are well defined the pattern was fitted at once, hence all the well-isolated lines in the pattern fix positions of lines in the group. they are all bound together by the lattice parameters. but then the two excessive lines must belong to other phase. it is concluded that this phase is r-phase. as the algorithm could not move the lines, it could minimize the sum of squares only using height, width and shapes of lines. the height is most significant factor then, and its adjustment subsequently influences the resulting austenite or martensite volume fraction. but this compensation is not perfect, which implies the loss of overall sum of the fractions. r-phase is usually found somewhere in between the austenite and martensite. if the interface is more or less of a planar aspect then from some angle the microfocused beam could be passing through the r-phase region rather perpendicularly (scans 11 and 66 in figure 6) and so the irradiated r-phase volume would be low. on the other hand it would be significantly higher if the beam was passing along the r-phase region (scans 36 and 96). in that case the sum of diffracting volumes evaluated from the presented fit would be most deficient (as explained above). when the overall sum of diffracting volumes was subtracted from virtual wire profile, pictures in figsin( 3)/6 [nm -1 ] 0.42 0.44 0.46 0.48 0.5 0.52 in te n si ty a .u . #10 4 0.5 1 1.5 2 2.5 3 3.5 4 4.5 data fit aust+mart fit aust fit mart lines mart lines aust (1-11) (110) (021) (111) (002) figure 7. central group of peaks proving the presence of r-phase. positions of austenitic, resp. martensitic lines are denoted by crosses of the same color as the fit of the given phase (austenite is denoted by blue, martensite by red color). green line stands for the sum of the former two fits, black dots are measured values. ure 8 were obtained. these pictures depict this difference relative to the virtual wire profile, i.e. (vvirt −vcalc)/vvirt. because the absorption correc11 martin dudr, daniel šimek, jan drahokoupil acta polytechnica ctu proceedings mm 0 2 4 6 8 10 m m 0 0.2 0.4 0.6 0.8 1 -1 -0.8 -0.6 -0.4 -0.2 0 0.2 0.4 0.6 0.8 1 mm 0 2 4 6 8 10 m m 0 0.2 0.4 0.6 0.8 1 -1 -0.8 -0.6 -0.4 -0.2 0 0.2 0.4 0.6 0.8 1 figure 8. relative difference of calculated diffracting volume from virtual wire profile with respect to wire profile (symbolically (vvirt − vcalc)/vvirt). on the left with fibre-texture correction, on the right with no-texture correction. tion was not carried out thoroughly, some excess of calculated diffracting volume was observed on the borders and a systematic decrease in the central part. still, some deviation from this behavior is well noticeable. notice that extreme values (±1) are nearly only on the wire border. notice also that the virtual wire profile is only illustrative and serves for visualization of "shape" of the sum of diffracting volumes. it is also possible to see that the regions of greatest decline (orange-red regions in figure 8, where r-phase is expected, especially in the figure on the left) are the same as the regions where the most austenite was observed (compare with figure 3). from this we conclude that in this case we are unable to distinguish between austenite and r-phase. 4. conclusion a simplified quantitative analysis of 2d diffraction patterns was carried out from transmission diffraction measurements of niti spring at european synchrotron radiation facility (esrf), grenoble. routines were created, which enabled automatic integration of 2d diffraction patterns and successive fitting of 1d patterns regarded as a two-phase mixture (whole-profile fitting based on analytical approach and levenbergmarquardt method of nonlinear least squares). with simplifying assumptions about texture it was possible to roughly estimate the ratio of the phases (austenite and martensite). martensite appears to be strongly prevalent in the sample. as the r-phase (which was present) was not taken into account the fit had troubles with convergence in those images where the r-phase occured significantly. these images show a loss of calculated overall diffracting volume. regions of higher austenite amount seem to be correlated with regions showing this loss, i.e. with regions where r-phase is expected too. on basis of presented results these two are not distinguishable. acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs16/245/ohk4/3t/14. references [1] šittner, landa, lukáš, novák. r-phase transformation phenomena in thermomechanically loaded niti polycrystals. mechanics of materials 38:475–492, 2006. [2] t. goryczka, h. morawiec. structure studies of the r-phase using x-ray diffraction methods. journal of alloys and compounds 367:137–141, 2004. [3] v.k. pecharsky, p.y. zavalij. fundamentals of powder diffraction and structural characterization of materials, 2nd edition, springer, usa, e-isbn: 978-0-387-09579-0, 2009. [4] fort wayne metals specification of nitinol#1 alloy http://www.fwmetals.com/services/ resource-library/nitinol-1/ [5] esrf official web page: id15a materials chemistry and materials engineering. [2016-10-14] , http://www.esrf.eu/usersandscience/experiments/ structmaterials/id15a. [6] h. p. stuwe,y. shimomura. gitterkonstanten der kubisch raumzentrierten phasen feti, coti, niti. zeitschrift fur metallkunde 51:180-181, 1960. [7] r. kužel. informace v práškovém difraktogramu a jeho zpracování. materials structure 10(1a): 18-19, 2003. 12 http://www.fwmetals.com/services/resource-library/nitinol-1/ http://www.fwmetals.com/services/resource-library/nitinol-1/ http://www.esrf.eu/usersandscience/experiments/structmaterials/id15a http://www.esrf.eu/usersandscience/experiments/structmaterials/id15a acta polytechnica ctu proceedings 9:1–5, 2017 1 introduction 2 material and methods 2.1 nitinol 2.2 quantitative phase analysis 2.3 experimental 2.4 data processing 3 results and discussion 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0024 acta polytechnica ctu proceedings 12:24–31, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app development of an electric bicycle for a sharing system in prague william deleenheera, ∗, lukáš jánešb, akshaya jayakumarb a kuleuven, faculty of industrial engineering, gebroeders de smetstraat 1, 9000 gent, belgium b czech technical university in prague, faculty of transportation sciences, department of vehicle technology. konviktská 20, 110 00 prague ∗ corresponding author: william.deleenheer@gmail.com abstract. by means of a development of an e-bike sharing system the electromobility project wants to provide an alternative way of comfortable transportation for students and staff of the ctu, primarily to commute between different campuses. the research for this project contains at least three different fields of study, namely electric vehicle and docking station development, intelligent transport systems and management and economics of transportation and telecommunication. after briefly stating general requirements for the sharing system, this paper focuses on the development of the electric bicycle. first an ideal bike design is defined. then necessary motor power and battery capacity are calculated by estimating characteristics of cycling in prague. a prototype was developed by converting a normal bicycle to an electric bicycle. being equipped with devices for e-bike monitoring, controlling and data recording for a post trip analysis, this prototype is also intended to have an educational value for future students in the project. results consist of an electrical bicycle configuration that matches the requirements and a sketch of an ideal e-bike for this project. keywords: e-bike sharing, requirements, e-bike design, motor power, battery capacity. 1. introduction the demand and supply of electric bicycles is a fastgrowing market. a market study on e-bike purchasing has increased for this year, mainly in the netherlands, denmark and germany. it is said that the market first concentrated on the netherlands and germany due to its overall size. figure 1. e-bike unit sales per 1000 inhabitants in 2015, data from ecf advocacy group’s publication[1]. but eventually the per head market demand increase in switzerland, austria, denmark proved to be relatively stronger. customers start to have innumerable options to combine any kind of bicycle with different sorts of batteries and motors at a wide range of prices. nevertheless, to compose a suitable e-bike for a sharing system is less easy. first of all, a responsibility that comes with owning a bicycle lies not with the end user anymore and thus the user has no incentive to care for it or maintain it. and second a sharing system takes into account many bicycles thus the cost of maintenance or theft as well as an initial investment of one e-bike should be as low as possible. furthermore, there are some requirements for a successful and economically viable sharing system, mostly regarding user experience, such as safety and comfort. this paper has tried to list these requirements and develop an electric bicycle accordingly. 2. general requirements for the sharing system determining the requirements for a viable e-bike sharing system for the ctu but also for prague in general has been carried out in three ways. in the first place, there was an existing project vision that originated both in the project goal as in earlier research. then more detailed requirements were defined by a research of existing examples in the academic world and in the market. finally, the requirements as for the station and software for an initial pilot project were summarized. 24 http://dx.doi.org/10.14311/app.2017.12.0024 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 development of an electric bicycle in prague 2.1. project vision the primary goal of this project is to provide a more efficient way of transportation for students and staff of the faculty of transportation science as a means to commute among different faculty buildings. the challenge exists in achieving this goal without losing comfort that comes with using a currently available public transport system. the concept of comfort in this context can be specified in four clear requirements: • minimal human physical effort, • protection against bad weather conditions, • possibility to transport certain cargo, • easy access to the system. a priority condition for all other requirements is obviously safety, all the more because cyclists are vulnerable road users and the bicycle infrastructure is still limited in prague. 2.2. market research market examples and previous experience with both conventional bike sharing schemes as well as a few existing e-bike sharing schemes such as go bike[2], bewegen.com[3], bonopark[4] and cycleushare[5] were researched to define components of a successful e-bike sharing system. another useful source was the experience of the institute of transportation and development policy (itdp)[6] with planning and exploitation of bike sharing systems. the following characteristics have been found repeatedly and are here briefly summarized: • integrated and theft protected e-bike components (motor, battery, chain, lights, wires, hid, automated lock); • ergonomic bicycle design (step-thru frame with adjustable saddle); • user friendly bicycle checkout and payment procedures; • high station density (10-16 stations for every square kilometre for conventional bikes); • many bicycles (10-30 bikes for every 1,000 residents for conventional bikes); • large coverage area (10 square kilometres for conventional bikes); • others: auxiliary (digital) lock, gps tracking, modular docks. discussable options are a use of high-end integrated components such as internal gears, belt drives or puncture free tires that reduce maintenance or theft costs but require higher initial investment. the university of brighton in uk conducted a similar e-bike research. the research aimed to understand user needs and to develop a system design necessary to start a fleet of e-bike program. the main outcome was a development of a "smart tracking system" which was implemented in a few e-cycles and proved successful. the system autonomously recorded and transmitted a bikes position, route, level of assistance required in real time and then sent it all in an online interface which was used for both research analysis and participant review. this proved to be an efficient solution for the fleet management and theft control. most of the users were daily commuters to office, the surveys revealed that most of them had found it easy to use and that it was something which had not disturbed or damaged their clothing much. 2.3. station and software requirements this project opted for a system in which the batteries have to be rented and plugged onto a bike by the user. this concept of battery charging and vending machine was first used in the "cycleushare project"[5]. the main advantage is that the station can have more batteries than bicycles and thus always has a fully charged battery for the user. in this concept all the batteries are stored in one charging box which protects the batteries from weather conditions and theft. this way the charging system and the bicycle docks can be made much cheaper compared to a more common way in which every dock needs a charger and the battery is integrated in the bicycle. another advantage is that the batteries, which mostly have a shorter life span than the bicycles themselves, can be maintained or replaced very easily. the biggest disadvantage is a decreased ease of use since the check in and check out procedure involves handling a battery. more experience is needed to determine an exact opportunity cost of this concept. the system is required to be an intelligent system. this includes necessity of mutual communication between the user, server, station, bicycle and battery. the tools to be used for this communication are isic student cards (rfid), gps tracking, rfid controlled locks, hid’s, system managing software and a web application. the main requirements for the software are based on the characteristics of sharing software like the one from "opensourcebikeshare"[7] and general requirements. crucial properties of the software interacting with the users via a web application are a map with real time data about station proximity and bicycle availability, instructions and help for users, feedback from users, reward system for a balanced e-bike distribution. additionally, it could have a routing tool if this is not integrated in the e-bicycle hid. the gui interacting with the administrators needs to provide data about system usage and provide a possibility to manage the fleet in real-time. this is possible via communication with rfid and gps tracking devices or if you like on docks and bicycles as well as user feedback. additionally, it will be connected to the battery charging machines which will provide data about battery state, such as number of cycles and usage patterns. 25 w. deleenheer, l. jáneš, a. jayakumar acta polytechnica ctu proceedings 3. electric bicycle research and design the above stated requirements have been applied on a development of a suitable electric bike. this involved researching an optimal bicycle design as well as a simulating power output and energy consumption during cycling to determine necessary motor and battery specifications. 3.1. bicycle design properties of bicycles in an average bike-sharing scheme are already a good approach to what a comfortable e-bike should look like. they are city-bikes with a step-thru frame, a rear rack for panniers or a basket mounted to the handlebars. with regards to cobblestone streets in prague it is desirable to have some form of suspension. e-bikes with the same purpose should have a similar design except the design also involves a decision about where to place a motor and battery. the requirement of a minimal human effort includes optimal handling and steering behaviour of the bike. the centre of a mass of the rider-bike system affects this behaviour. in fact, the act of cycling is known to be a fairly complex mechanism. in addition to the rider’s skill and gyroscopic forces, there are, acting on the wheel, the centre of gravity lowering torque and castoring forces[8]. in the context of this project we can simplify it by saying that keeping a bicycle from falling comes down to keeping the centre of mass of the rider-bike system over its wheels. at very low speeds or when coming to a halt, the bike is easier to handle with a low centre of mass because thanks to the the leverage, the rider also needs to apply relatively small forces on the handlebars to keep the bike upright. from a certain minimum forward speed, it is easier to steer a bike with a higher centre of mass. f for the same reason it is easier to balance a tall broomstick on the end than a short pencil. the broomstick is a slower inverted pendulum for which one has more time to bring it back to balance[9]. on the other hand, as it is the case for normal bicycles, the mass of the rider still has the most impact on the overall centre of mass of the coupled electric bike – rider system. thus for a placement of a battery and motor it is sufficient to look at handling behaviour at very slow speeds or standstill. furthermore, taking in consideration the idea of transporting certain cargo on the rear rack it is preferable to have the battery weight inside the frame for a better mass distribution, for a minimal inertia high in the rear and thus for a safe and comfortable ride. the decision to choose a mid-drive motor contributes to this logic but is also a consequence of reasons stated in the next paragraph. disadvantages of hub motors are an increased unsprang weight and a more difficult maintenance. apart from an extra mass added to the bicycle, the motor and battery need extra space. one idea on this matter is to use smaller wheels (24 to 18 inch) which are already often used in cargo bikes for this reason. smaller wheels also lower the overall centre of the mass and enable higher acceleration amongst other advantages[10]. one disadvantage is a decreased cushioning. a compromise could be only to have a smaller rear wheel. figure 2. e-bike design: sketch by the writer, scale 1/10. to visualize the previously outlined ideas, the above draft of a possible e-bike design was sketched, keeping in mind low cost, easy fabrication for example with standard aluminum tubes. this draft could be a base for a more advanced (e.g. with integrated components) and attractive design that can be validated with a strength analysis. a last point to look at is a protection against wind, rain and sunlight. there are some solutions on the market to solve this problem, such as detachable shields or poncho’s. in case a full integration of this feature is opted, it is recognizable that the lower the rider sits to the ground the easier it is to design a cover/shield that protects a full body of the rider from the impact of precipitation and wind owing to its forward movement. 3.2. motor the main specification of an electric bike motor is its power. on the market the motors are distinguished by an electrical power they can infinitely continuously handle, for example "a 250 watt motor". this figue is misleading because of two reasons: the input power the motor uses is determined by the voltage of the battery multiplied by the current the controller draws from the battery (e.g. 36v × 15a = 540 watt). the motor will mostly be built for a certain voltage (e.g. a 36v 250 watt motor with a 36v battery) so the maximum power will be proportional to a current limit in the controller. secondly the electrical power is not equal to an output power on the axis of the motor because of internal losses. the latter one is the power assisting the rider in giving a bicycle a forward motion. the average efficiency η of most motors on the market is around 80% at the optimal rpm so 350 watts into the controller will result in 280 watts of a driving power[11]. 26 vol. 12/2017 development of an electric bicycle in prague minimal human effort and a possibility to transport certain cargo are the two initial requirements which come in to play when determining required motor power for this project. in other words, it can be stated that cycling should not be experienced as an exercise and the user can go from a to b in normal clothing without sweating (disregarding the influence of abnormal values for humidity and ambient temperature). the question is which human output power corresponds to this requirement. professional cyclists can sustain 400 watts of leg power for an hour while for a person of an average health and fitness this would be 100 to 200 watts. these values apply for exercise till exhaustion[12]. for a comfortable commuting a continuous human power output of 50 watts is decided to be a reasonable maximum. this amount is comparable to the power needed for a person of 80kg to walk 1 m/s[13]. the last steps in defining a necessary motor power consist first of all of specifying a required driving power for cycling as a function of speed [km/h] and uphill slope [%] and last of simulating required performances for cycling in prague. at a constant speed the driver and motor combined have to overcome three forces: (1.) rolling resistance (frol), frol = mtot × g × cr where: mtot = mass of the rider-bike system, g = gravitational acceleration (9.81 m/s2) and cr = rolling resistance coefficient (±0.006 for normal tire) (2.) aerodynamic resistance (faer ), faer = 1 2 × ρ × v2 × cw × af where: ρ = air density (±1.20 kg/m3 for 20°c and 1 atm), v = bicycle speed [m/s], cw = air resistance coefficient (±1.1 for upright cyclist plus touring bike) and af = frontal surface area (0.6 m2 for upright cyclist) (3.) inclination resistance (fincl), fincl = mtot × g × sinα where: mtot = mass of the rider-bike system, g = gravitational acceleration (9.81 m/s2) and α = is the inclination grade in degrees then ftot = frol + faer + fincl pdriving,tot = ftot × v assumption for mtot = mdriver + mloadedbicycle = 75 + 35 = 110kg. all assumptions were made with a rather inefficient but still realistic set up in mind. the chosen parameter values are the result of a comparison of [12] and [14]. on figure 3: power input during acceleration => from 0 to 25km/h in 3 seconds => f = m×(v2−v1)/t. figure 3. driving power for speeds up to 50km/h and slopes up to 20%. simulating cycling in prague is a rather subjective process but as for a speed it is relevant to take 25km/h which is the maximum speed to which a motor can assist an electric bike by eu legislation. it is known that prague has uphill slopes up to 15% in the form of short streets. it is preferable these can be done with comfort and without the motor failing. with the above two values as a maximum the necessary driving power will be 1300w. eu legislation restricts a rated power to a maximum of 250w. in fact this stands for a 36v motor with a 15a controller which means this motor will be able to handle 540w although not continuously. in the latter case the output power will average around 430w which is far below a defined required driving power. as long as continuous comfort is more important for this project than a continuous speed, it is also useful to define a minimum speed. it was arbitrarily chosen to take 10km/h, which is still a double of an average walking speed, as a minimum speed to go uphill. now the maximum necessary total driving power is 472w. providing the rider takes 50w on his behalf, the required motor input power becomes (472 − 50) × 100/80 = 527.5w. it can be recognised that a motor rated as 250w is still insufficient, especially because at low speeds motor efficiency will be lower than 80% and the total mass might be higher. there is a clear preference of a mid-drive motor to a wheel hub motor or a direct drive for implementation in a hilly city like prague since a mid-drive motor can still work in (or closer to) the optimal rpm range at low speeds such as 10km/h provided a good use of the bicycle gears. present extensive e-bike testing and user experience will tell if this was a good simulation. 3.3. battery the battery is a fuel tank for an electric vehicle(ev). the required amount of energy it must store depends on energy consumed in a trip with a certain minimal range or minimal time. a range of 15 to 20km is 27 w. deleenheer, l. jáneš, a. jayakumar acta polytechnica ctu proceedings v [km/h] avgup hill [%] distance [km] #accel etot,el [wh] +human etot,el [wh] throttle mode 15 0 20 1 9.6 93.0 20 0 20 1 67.8 130.3 25 0 20 1 128.3 178.3 25 0 40 1 255.9 355.9 25 0 50 1 319.7 444.7 25 2 5 1 70.1 82.6 25 2 5 10 76.7 89.2 25 2 5 20 84.1 96.6 25 2 10 1 139.5 164.5 25 2 15 1 208.8 246.3 25 2 20 1 278.2 328.2 25 2 30 1 416.9 491.9 20 2 20 1 217.7 280.2 10 5 5 1 79.0 110.2 15 5 5 1 96.2 117.0 20 5 5 1 110.9 126.5 25 5 5 1 126.2 138.7 10 10 1 1 34.5 40.7 10 10 5 1 171.8 203.1 15 10 1 1 38.0 42.2 20 10 1 1 41.1 44.2 20 10 5 1 203.7 219.3 25 10 1 1 44.4 46.9 table 1. estimated energy consumption within 23 different trips. more than enough to cover an area that includes the center of prague and the majority of ctu buildings. the electrical energy e consumed is equal to the time integral of electrical power p(t). if p = constant for certain time interval ∆t, following expression applies: eelectric = pelect × t = η × pdriv,mot × t assuming the same values for the rider-bike system as in the previous section the energy consumption in over 230 trips was simulated by setting discrete values for the following variables: velocity [km/h] (10,15,20,25), average uphill slope [%] (0, 2, 5), distance [km] (1,3,5,10,15,20) and amount of accelerations at 1.388m/s2 (1,5,10). the table below (table ??) shows some of the more relevant possible trips including scenarios that did not fall into the scope of the simulation. the values in yellow show an effect on power and energy of changing this particular variable. the values in orange show trips with an energy consumption of more than 200w in the throttle mode. the values in green show trips which are more typical for a bike sharing system ( <5km & <0.5h). the values in red show trips that require more than 650w of an input power. taking into account that a deep discharge is to be avoided when trying to maximize a battery life span, it can be stated that 150 to 200wh of a stored energy will be more than enough in the context of this project. practically this could be a 36v × 6.6ah or a 25v × 8ah. the main specifications of a battery pack are voltage [v] and amp-hours [ah]. the electrical energy storable in the battery pack can be estimated with this theorem: eelectric = v (t) × i(t) × t where v (t) = rated voltage of the battery pack [v] and i(t) × t = current discharge over time [ah]. for the 36v motor from the previous section and 300wh, equation (7) gives a 36v × 8.3ah (theoretical) battery. in practice, a battery pack is configured by a spot welding a number of cells in series and parallel e.g. 3.6v × 2200mah cells in a 4parallel10serie configuration makes a 36v × 8.8ah battery pack. the number one chemistry for these cells is lithium-ion e.g. lifepo4 (ifr), licoo2 (icr), limn2o4(imr), linicoalo2 (inr), each with its own benefits and downsides. when optimizing this project, it will be highly relevant to compare these chemistries. for the latter examples it can be simplified by stating li-manganese are especially safe, li-iron-phosphate are the most economical (expensive but a high life span) and li-cobalt have the highest specific energy but they need a good bms to be safe[15]. another important property of a cell is its maximum continuous discharge rate. this rate is defined by its c-rate which expresses the time in which a battery can be discharged compared to its capacity in amp-hours, e.g. a 2c battery cell with 2200mah can be discharged at 4400ma in 30 minutes[16, 17]. if the motor controller draws equal or more amps from the battery pack than its configuration can handle, this will result in a decreased capacity and probably also a decreased life span. in other words: if the e-bike is equipped with a 15amp controller, a 2c rated battery will need to have at least 7.5ah but preferably 10ah. thus, it can be stated that for this e-bike sharing project it is preferable to have a cell chemistry with a high c-rate since enough power is more important than a high capacity depending on what is an average distance between the stations and the characteristics of an average trip. a standard size for these cells is 18650 (18mm diameter, 65mm long, round shape) and their weight is a little less than 50g. a charging rate is mostly between 0.5c and 1c e.g. a 8.8ah battery pack charged at 2a will take about 5-6 hours to be fully charged. note that it will take about as much time to charge a battery from 0 to 80% as from 80% to 100%. it was the case in the "cycleushare" project that the batteries were made available by the system from the moment they were 75% charged which increased efficiency but still a reasonable range[18]. 28 vol. 12/2017 development of an electric bicycle in prague 3.4. electronics just like the system itself is required to be intelligent, so much is the e-bike itself. via a gui on the handlebars the rider can be informed about multiple variables as speed, trip distance, trip time, instantaneous power input and battery capacity, etc. more advanced devices as the cycle analyst will also provide information about wh or ah consumed, estimated wh left (instead of the typical battery icon) and an amount of human power amongst other variables. if a gps tracker is built in and a logging device is installed on the bike it is possible to carry out full post trip analyses which are interesting for further research in this project. using a mid-drive motor, it is preferable to have a gear sensor that cuts off motor power momentarily when the user shifts gears, thereby preventing too much stress on the chain and the gearing system which causes rough gearing and can cause damage as well. a better but more expensive solution would be a continuously variable transmission (cvt). the motor is mostly controlled with hall sensors. there are three main ways the rider can get assistance from the motor: throttle mode (current flow is proportional to angle of (thumb) throttle), cadence based pedal assist (current flow is proportional to crank rotation speed), torque based pedal assist (current flow is proportional to amount of torque exerted on the chain or the pedals). from these three ways the latter feels the most intuitive while the cadence based pas is much cheaper and thus more common. another advantage of torque proportional assistance is an increased battery range[19]. to achieve the proclaimed 50w of human power at all speeds and slopes a torque sensor would be ideal. in practice this can be done with a chain tension sensor or a strain gauge on the cranks or pedal axles. usually the hid offers a chance to set different pas levels so the user can choose how much proportional assistance he or she requires from the motor. most battery packs on the market will have a battery management system (bms) which is a circuitry to keep individual cell voltages balanced during charging and discharging. furthermore, it protects the cells from different circumstances such as high temperatures, deep discharge and overcharging. the quality and intelligence of this bms will have a high impact on the life span of a battery pack and depending on the chosen cell chemistry the bms is essential for safety, more specifically for preventing fires or explosions. 4. e-bike building, testing and results 4.1. choosing a configuration due to a low initial budget it was decided to convert a normal bicycle to an e-bicycle with a commercially available mid-drive kit and a downtube battery. because this battery requires presence of water bottle cage mounting holes it was more convenient to buy a male framed bicycle (see figure 4). figure 4. photo of an assembled e-bike. the configuration of the research bike consisted of a 350w rated mid-drive motor, meaning a 36v motor with 18a controller, i.e. a peak input power of 648w. the battery pack has a capacity of 36v × 8.8ah which is a 4p10s configuration of samsung icr1865022p cells (3.62v × 2150mah). samsung claims these cells have a maximum continuous discharge rate of 3c which means 26.4a. in fact, discharging at 3c the battery will already put quite some stress on the cells and the capacity will be lower due to heat losses as can be seen on discharge curves of this particular cell[7]. 4.2. tests and results firstly, the balancing of the e-bike was found very easy. one indicator for this is that it was very easy to balance the bike without hands on the handlebar. the bike enabled flexible maneuvering. this proves the chosen placement of battery and motor to be correct. secondly acceleration, top speed in the throttle mode was up to 34 km/h on a flat in an upright position. the accelerations perceived a big difference with normal cycling. the ease of getting up to speed was so easy both with the pedal assistance and the throttle mode is similar to riding a scooter or a small motorcycle. the motor did tend to heat up quite quickly after riding the bike several minutes in the highest pas levels (4 and 5). this shows that the set up with 36v and 18a the limit approaches the level a motor with these 62 dimensions can handle. the same motor is rated "250w" with a 15a current limit, in this case the motor will probably not heat up as much, combined with less stresses on the plastic gears inside. it might therefore have a longer lifespan with 15a maximum. riding in prague with this e-bike also revealed possible dangers, mostly concerning the higher speeds achievable. being very vulnerable on an e-bike speeds tended to drop when riding among car traffic. 29 w. deleenheer, l. jáneš, a. jayakumar acta polytechnica ctu proceedings another inconvenience is that not being able to go to a lower gear at low speeds or standstill feels rather unhandy regarding smooth accelerations. this is not any different with normal bikes but with this e-bike there was a tendency to prevent shifting too much and staying in a high gear for higher speeds. because the brakes cut off the motor power there is no need to go down in gears when slowing down and coming to a stop. although the electric motor can handle starting in the highest gear, the acceleration at lower gears is more logical, but shifting to a lower gear at standstill causes much stress on the gearing system and is perceived unsmooth. 5. experiments 5.1. first experiment for last three years two experiments have been carried out with the e-bike for transportation among the three buildings of ctu. these two experiments have been performed by students of the project electromobility. the bikes for the experiments were borrowed with similar or lower configuration as the bicycle, which was constructed. the first experiment covered the connections between the three buildings of ctu. in this project only students of ctu were included. in every buildings there was a spot for borrowing an e-bike, where the students borrowed the e-bike and rode to another building of ctu, where they returned the e-bike. during the trips the students measured parameters of the trip by mobile application and sent data to the folder for processing. in 15 days there were made about 50 rides and every ride proved that transport by e-bike among buildings is faster than other types of transport, such as automobile or public transport[20]. 5.2. second experiment the second experiment was made in 2014, and it examined advantages and parameters of riding the e-bike to the center of prague from an outskirt of the city. for this experiment typical trips for riding to center of prague were chosen, the center was selected as the building of ctu horská. there were 10 trips and every measuring showed that every trip is faster by e-bike opposite to the car or public transport. only one trip, the trip of about 10 kilometers, was faster by metro than e-bike[21]. 5.3. experiments conclusion the e-bike constructed for this thesis and presented in this article, was tested only for the driver’s comfort and for technical parameters, but not for measuring parameters and comparison with another type of transport. the previous two experiments were only for testing e-bikes as a means of transport for students among the buildings of ctu or for transporting students from home to school. it should start as a private public transport for students in "campuses", with private spots for rental and private system of borrowing and security. in the future, it would be beneficial to add new places, such as the central station, where it can be extended to general public which will help to make the system more popular among general population and tourists. 6. conclusion it is clear that e-bike sharing is a useful technology which can be a transportation method made ready to fit to a wide range of possible users. however, the small portion of real world examples show the concept still faces some difficulties, most of which being financial or organizational ones. thus, it is important to plan implementation and exploitation thoroughly with all stakeholders involved in the debate with a strong business plan laid on the table. the research e-bike, equipped with the cycle analyst and logger can be used to acquire accurate data about energy consumption and other trip characteristics when e-cycling in prague meanwhile experimenting with different types of bicycles, installing wind and rain shields or full covers to see their effect on energy consumption and comfort. as a result, the developed e-bike can be optimized further and fabricated. specific recommendations for further successful realization of the e-bike sharing project: • set up a strong business plan by means of a user survey and investment analysis and specific action plans. • further visualize the final result by implementing a (3d) design of both station and e-bike. • cooperate with other institutions and faculties. • set up an organization which will manage this interdisciplinary project and represent it. these 4 recommendations will significantly help with the challenge of an increasing student interest and involvement and increasing interest from sponsors, advertisement companies, university management and city municipalities. according to the market studies prague manages to keep up with the trends and efficient modes of transport. e-bike sharing could be then very soon visualized in prague at ctu. references [1] "european bicycle market analysis 2015", advocacy means sales, european cyclists’ federation. edited by: alexander scholz, published on 25.08.2016. [2] "gobike", gobike. [online], available: http://gobike.com/solution/the-bike/, [geopend 10 may 2016]. [3] "bewegen", bewegen technologies inc. [online], available: http://bewegen.com/our-system/#top, [geopend 10 may 2016]. 30 http://gobike.com/solution/the-bike/ http://bewegen.com/our-system/#top vol. 12/2017 development of an electric bicycle in prague [4] "bonopark", bonopark. [online], available: http://www.bonopark.es/en, [geopend 10 may 2016]. [5] "cycleushare", university of tennessee. [online], available: http://www.cycleushare.com/, [geopend 10 may 2016]. [6] "itdp", the institution of transportation and development policy. [online], available: https://www.itdp.org/wp-content/uploads/2014/ 07/itdp-bike-share-planning-guide-1.pdf, [geopend 10 may 2016]. [7] "open source bike share". [online], available: http://opensourcebikeshare.com/, [geopend 25 may 2016]. [8] d. e. jones. the stability of the bicycle. physics today. pp. 34-40, 1970. [9] j. m. j. p. a. r., a. j. d. g. kooijman. a bicycle can be self-stable without gyroscopic or caster effects. science, no. 332, pp. 339-342, 2011. [10] t. hadland. small wheels for adult bicycles. 25 june 2012. [online], available: https://hadland.wordpress.com/2012/06/25/ small-wheels-for-adult-bicycles/, [geopend 25 may 2016]. [11] j. lemire-elmore. ebiks.ca. grin technologies. [online], available: http://www.ebikes.ca/tools/simulator.html, [geopend 28 april 2016]. [12] m. ikonen. ip-zev. 2 april 2011. [online], available: http://www.ip-zev.gr/files/teaching/t4-2_human_ power.pdf, [geopend 28 april 2016]. [13] d. j. farris, g. s. sawicki. ncbi. pmc, 25 may 2011. [online], available: http://www.ncbi.nlm.nih.gov/ pmc/articles/pmc3223624/table/rsif20110182tb1/, [geopend 28 april 2016]. [14] "coefficient of friction, rolling resistance and aerodynamics", tribology-abc. [online], available: http://www.tribology-abc.com/abc/cof.htm# rolweerstand, [geopend 30 may 2016]. [15] "battery chemistry finally explained", battery bro, 19 february 2015. [online], available: https://batterybro.com/blogs/ 18650-wholesale-battery-reviews/ 18880255-battery-chemistry-finally-explained, [geopend 26 may 2016]. [16] "batteries", grin technologies. [online], available: http://www.ebikes.ca/learn/batteries.html, [geopend 26 may 2016]. [17] "bu-402: what is c-rate?", battery university, 29 february 2016. [online], available: http://batteryuniversity.com/learn/article/ what_is_the_c_rate, [geopend 27 may 2016]. [18] c. r. cherry, s. ji, l. d. han, d. a. jordan. electric bike sharing: simulation of user demand and system. journal of cleaner production. vol. 85, pp. 250-257, 2014. [19] "torque sensors on electric bikes", ectricbike.com, 18 october 2012. [online], available: http://www.electricbike.com/torque-sensors/, [geopend 27 june 2016]. [20] m. lapáčková. bike sharing na fakultě dopravní. praha, 2014, diplomová práce. čvut fd. [21] l. jáneš. využití elektrokol v praze z hlediska bezpečnosti a členitosti terénu. praha, 2015, bakalářská práce. čvut fd. 31 http://www.bonopark.es/en http://www.cycleushare.com/ https://www.itdp.org/wp-content/uploads/2014/07/itdp-bike-share-planning-guide-1.pdf https://www.itdp.org/wp-content/uploads/2014/07/itdp-bike-share-planning-guide-1.pdf http://opensourcebikeshare.com/ https://hadland.wordpress.com/2012/06/25/small-wheels-for-adult-bicycles/ https://hadland.wordpress.com/2012/06/25/small-wheels-for-adult-bicycles/ http://www.ebikes.ca/tools/simulator.html http://www.ip-zev.gr/files/teaching/t4-2_human_power.pdf http://www.ip-zev.gr/files/teaching/t4-2_human_power.pdf http://www.ncbi.nlm.nih.gov/pmc/articles/pmc3223624/table/rsif20110182tb1/ http://www.ncbi.nlm.nih.gov/pmc/articles/pmc3223624/table/rsif20110182tb1/ http://www.tribology-abc.com/abc/cof.htm#rolweerstand http://www.tribology-abc.com/abc/cof.htm#rolweerstand https://batterybro.com/blogs/18650-wholesale-battery-reviews/18880255-battery-chemistry-finally-explained https://batterybro.com/blogs/18650-wholesale-battery-reviews/18880255-battery-chemistry-finally-explained https://batterybro.com/blogs/18650-wholesale-battery-reviews/18880255-battery-chemistry-finally-explained http://www.ebikes.ca/learn/batteries.html http://batteryuniversity.com/learn/article/what_is_the_c_rate http://batteryuniversity.com/learn/article/what_is_the_c_rate http://www.electricbike.com/torque-sensors/ acta polytechnica ctu proceedings 12:24–31, 2017 1 introduction 2 general requirements for the sharing system 2.1 project vision 2.2 market research 2.3 station and software requirements 3 electric bicycle research and design 3.1 bicycle design 3.2 motor 3.3 battery 3.4 electronics 4 e-bike building, testing and results 4.1 choosing a configuration 4.2 tests and results 5 experiments 5.1 first experiment 5.2 second experiment 5.3 experiments conclusion 6 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2018.17.0006 acta polytechnica ctu proceedings 17:6–9, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app investigation of surface macroscopic residual stresses of cutting ceramics after mechanical machining jakub němeček∗, jiří čapek, nikolaj ganev, kamil kolařík department of solid state engineering, faculty of nuclear science and physical engineering, czech technical university in prague ∗ corresponding author: jakub.nemecek@fjfi.cvut.cz abstract. currently, the extensive research in the field of cutting ceramics is conducted and there are efforts to replace cemented carbides by these materials. however, the problem is to improve the production of ideal compact samples. therefore, the influence of mechanical machining technologies on the values of macroscopic residual stresses was investigated by x-ray diffraction. the resulting values were discussed depending on machining parameters and surface structure of the studied samples. keywords: cutting ceramics; x-ray diffraction; residual stress; machining technology. 1. introduction a development of cutting ceramics has begun at the beginning of 20th century. nowadays, there is an attempt to optimize the production process and machining parameters depending on the real structure of ceramic tiles [1–4]. in cooperation with moscow state technological university stankin (mstu stankin), which provides preparation of samples, the macroscopic residual stresses were analysed by x-ray diffraction. analysed samples were prepared from different materials such as trigonal al2o3 (samples 2.x), hexagonal si3n4 (samples 3.x), hexagonal sic (samples 4.x), and tetragonal zro2 (samples 5.x). they were finished with different mechanical surface treatments, specifically grinding by diamond disc (samples x.1−x.3) and air abrasive treatment (aat, samples x.9−x.11), see table 1 and 2. samples of first series (1.x) were analysed in [5]. production of samples differed for individual materials. samples of al2o3, sic and zro2 were made by so called standard sintering, when they were pressed at low temperature and then put into a furnace, where the bonds in material were created. samples of nitride ceramics were made by cold sintering, but they were put into compressible ampoules which the hydrostatic pressure exerted during the sintering process. this has resulted in high material densities, see [6]. sample material 2.x al2o3 3.x si3n4 4.x sic 5.x zro2 table 1. materials of individual series of samples. sample technology x.1−x.3 grinding by diamond disc x.9−x.11 air abrasive treatment table 2. machining technologies of individual series of samples. 2. experiment the influence of machining parameters on the surface values of macroscopic residual stresses was studied by x-ray diffraction. for grinding by diamond disc, the impact of transverse displacement on the lift f d was studied, and for the aat, where the sample rotated, the influence of loading time t was analysed (see tables 3 and 4). values of residual stresses were determined in two orthogonal directions l (direction of grinding) and t (direction of transverse displacement). the tensometric analysis of macroscopic residual stresses had different parameters for individual series of samples. for al2o3 diffraction line {0210} was analysed by manganese anode on diffraction angle 2θ = 146o and x-ray diffraction constants s1 = −0.61 tpa−1, 12s2 = 3.18 tpa −1, for si3n4 diffraction line {201} by chromium anode on diffraction angle 2θ = 63o in ψ-geometry and x-ray diffraction constants s1 = −0.87 tpa−1, 12s2 = 3.87 tpa −1, for sic diffraction line {028} by manganese anode on diffraction angle 2θ = 149.5o and x-ray diffraction constants s1 = −0.41 tpa−1, 12s2 = 2.74 tpa −1 and for zro2 diffraction line {220} by chromium anode on diffraction angle 2θ = 128o and x-ray diffraction constants s1 = −1.45 tpa−1, 12s2 = 6.42 tpa −1. for more details see [6]. 3. result and discussion at first, all samples were machined by diamond disc and some of them subsequently machined by air abra6 http://dx.doi.org/10.14311/app.2018.17.0006 http://ojs.cvut.cz/ojs/index.php/app vol. 17/2018 investigation of surface macroscopic residual stresses of cutting ceramics sample f d ,[mm/displ.] x.1 0.5 x.2 1.0 x.3 1.5 table 3. parameters of grinding by diamond disc. sample t,[s] x.9 15 x.10 30 x.11 60 table 4. parameters of aat. sive treatment. depending on the structure, various effects of machining technology on the surface structure were found. for trigonal al2o3, the values of residual stresses have expecting character, i.e. while after grinding by diamond disc in direction l, values of compressive residual stresses decrease with increasing f d, in direction t, the compressive residual stresses increase (see figure 1). this effect is caused by moving the cutting tool in the l direction, when the so-called stretching of the subsurface areas occurs due to the mechanical interaction of the cutter blades with the material. in the t direction, the higher feed rate leads to higher stresses. in the case of air abrasive machining, it is evident that with the increasing time of machining, the values of the compressive residual stresses increase while the isotropy of the values of the macroscopic residual stresses in the l and t directions are increased. this is caused by the rotational movement of the sample during machining. with longer time of machining the depth of the infuenced surface layer of material increases. to prove that is necessary to determine precisely the depth gradient of macroscopic residual stresses by forceless polishing. in the case of si3n4 (see figure 2), the residual stresses were analysed by different experimental geometry due to overlap of diffraction lines with diffraction angles above 110o 2θ. it means the resulting values of residual stresses are from thinner surface (ca. 4× thinner). an accuracy of determination of position 2θ is about 8, 5× smaller than for diffraction lines with higher angles 2θ. however, the relative dependence for determining the effect of mechanical machining on surface residual stresses is indicative. after diamond wheel grinding, the compressive residual stresses increase in the both directions with increasing f d. this is caused by hexagonal close packed structure of si3n4 and its resistance to machining. as it has been shown in [7], machining materials with the hexagonal close packed lattice can cause an increase of compressive residual stresses in the grinding direction. this effect figure 1. values of macroscopic residual stresses after grinding by diamond disc (samples x.1-x.3) and aat (samples x.9-x.11) of al2o3. figure 2. values of macroscopic residual stresses after grinding by diamond disc (samples x.1-x.3) and aat (samples x.9-x.11) of si3n4. is induced by the reaction of the structure to a force which resists plastic deformation. after air abrasive treatment of si3n4 samples, the effect of the previous machining technology can be seen. at a short machining time, the macroscopic residual stresses have a relatively large anisotropy and, with increasing machining time, this anisotropy disappears. at the same time, it can be seen that with the machining time, the values of the surface compressive residual stresses increase, which is again caused by the hexagonal structure. compressive stresses dropped for sample 3.11 probably by sliding along the grain boundaries (there was an increase in crystallite size, see [8]). for hexagonal sic, it can be seen that in case of grinding with a diamond disc in the direction of grinding (direction l), the values of the compressive residual stresses increase and in the direction perpendicular (direction t), the tensile residual stresses decrease. this dependence is inconsistent with expectations, since in general grinding causes tensile stresses in the direction of grinding and in the perpendicular direction the compressive residual stresses. it is caused 7 j. nemecek, j. capek, n. ganev, k. kolarik acta polytechnica ctu proceedings figure 3. values of macroscopic residual stresses after grinding by diamond disc (samples x.1-x.3) and aat (samples x.9-x.11) of sic. figure 4. values of macroscopic residual stresses after grinding by diamond disc (samples x.1-x.3) and aat (samples x.9-x.11) of zro2. again by hexagonal close packed structure of sic, see [7]. values of macroscopic residual stresses after air abrasive treatment of sic samples show very high experimental errors. these are due to the large dimensions of crystallites. however, it is evident that in the l direction, the values of the compressive residual stresses decrease, even tensile stresses increase, and in the t direction the compressive stresses on the contrary increase. this is caused again by the influence of the hexagonal lattice and its resistance to plastic deformation (see figure 3). after grinding by diamond disc of zro2, it is evident that the values of the compressive macroscopic residual stresses decrease with the increasing f d. this is due to the thermal overload of the surface. the thermal conductivity coefficient for zro2 is approximately one-tenth compared to other studied materials (λ = 2, 7 wm−1k−1 for zro2, other ceramics have a temperature coefficient λ = 30 wm−1k−1 and more). for air abrasive treatment, it is obvious that due to the rotational movement of the zro2 samples, an isotropic distribution of macroscopic residual stresses figure 5. surface defects of sic cutting ceramic after grinding by diamond disc (5a, 5b) and aat (5c, 5d). has occurred, the values in the l and t directions are same within the error (see figure 4). for samples from sic the analysis of surface with sem was made. the purpose was to investigate a presence of cracks, which may be caused by large tensile residual stresses. however, it is evident from 8 vol. 17/2018 investigation of surface macroscopic residual stresses of cutting ceramics figure 5 that no cracks were found. but the images show, alongside smooth cutting traces, surface defects caused by grinding, such as various holes or ejected grains. the morphology of surface is significantly changed since the cutting traces of diamond wheel disappeared and instead can be seen deep craters after the impact of individual grains of abrasive. 4. conclusions the analysis of macroscopic residual stresses of surfaces of four different ceramic materials was performed by x-ray diffraction and the effect of mechanical machining on the surface structure was determined. different reactions of hexagonal structures on mechanical machining were observed. it was found that zro2 due to poor thermal conductivity had to be machined at lower speeds or pressures, otherwise the heat load of the structure and the tensile residual stresses would increase. acknowledgements this work was supported by grant student grant competition ctu no. sgs16/246/ohk4/3t/14 and grant agency of the czech republic (no. 14-36566g). references [1] a. humár. materiály pro řezné nástroje. first printing. mm publishing, 2008. [2] v. v. kuzin, et al. correlation of diamond grinding regimes with al2o3-tic-ceramic surface condition. refractories and industrial ceramics 57(1):520–525, 2017. doi:https://doi.org/10.1007/s11148-017-0016-0. [3] v. v. kuzin, et al. correlation of diamond grinding regimes with sisic-ceramic surface condition. refractories and industrial ceramics 58(2):214–219, 2017. doi:https://doi.org/10.1007/s11148-017-0108-x. [4] v. v. kuzin, et al. correlation of diamond grinding regime with surface condition of ceramic based on zirconium dioxide. refractories and industrial ceramics 57(1):625–630, 2017. doi:https://doi.org/10.1007/s11148-017-0035-x. [5] j. němeček, k. kolařík, j. čapek, g. n. influence of machining technologies on values of residual stresses of oxide cutting ceramics. acta polytechnica ctu proceedings (9):13–18, 2017. doi:https://doi.org/10.14311/app.2017.9.0013. [6] j. němeček. difrakcní studium vybraných charakteristik řezné keramiky, prague 2017. master thesis (ing.). czech technical university in prague, faculty of nuclear science and physical engineering, department of solid state engineering, 2017-06-16. [7] k. t. ramesh. effects of high rates of loading on the deformation behavior and failure mechanisms of hexagonal close-packed metals and alloys. metallurgical and materials transastions a 33:927–935, 2002. doi:https://doi.org/10.1007/s11661-002-0162-x. [8] z. yin, et al. friction and wear behaviors of al2o3/tic micro-nano-composite ceramic sliding against metals and hard materials. ceramic international 42:1982–1989, 2016. doi:https://doi.org/10.1016/j.ceramint.2015.10.001. 9 http://dx.doi.org/https://doi.org/10.1007/s11148-017-0016-0 http://dx.doi.org/https://doi.org/10.1007/s11148-017-0108-x http://dx.doi.org/https://doi.org/10.1007/s11148-017-0035-x http://dx.doi.org/https://doi.org/10.14311/app.2017.9.0013 http://dx.doi.org/https://doi.org/10.1007/s11661-002-0162-x http://dx.doi.org/https://doi.org/10.1016/j.ceramint.2015.10.001 acta polytechnica ctu proceedings 17:6–9, 2018 1 introduction 2 experiment 3 result and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0029 acta polytechnica ctu proceedings 5:29–37, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app assets of model metro station and their criticality tomáš kertis∗, dana procházková czech technical university in prague, faculty of transportation sciences, praha, czech republic ∗ corresponding author: kertitom@fd.cvut.cz abstract. the paper deals with the identification of protected assets in the model metro station, which is a part of critical infrastructure in praha; and with the identification of their criticality. a metro station is used by a lot of people and it involves many expensive components and devices with interdependences which form a system of systems. the system, its components and fittings are very vulnerable to damages from disasters of many kinds. the first step of ensuring metro station safety is the identification of basic metro assets which are important for safe metro operation and for human (employee and passengers) security and environment safety; and the other is determination of criticalities of considered assets with taking into account all relevant sources of hazards. these facts are input data for further processes which are important for safe metro operation under the different conditions. keywords: safety, security, metro station, assets, disasters, criticality. 1. introduction a metro is very important means of transportation in cities. therefore, to its safety it is pursued special attention, and it is part of critical infrastructure, to which it is now concentrated special attention because especially at critical situation its status predetermines the capability of human society to ensure human survival [1]. metro is composed from fix and movable parts; the fix ones are structure element, stations and tunnels with many fittings; the movable ones are trains [2]. the paper takes into account the metro stations. each metro station is equipped by many technical devices and fittings that ensure metro operation and comfort space for passengers. from safety reasons it also involves support elements for safety operation of station; they are for instance construction design, operation devices and systems, control systems, protection systems, electronic devices like lifts, escalators, air conditioning, trains and systems for train control, communication systems, staff etc. mentioned elements including the humans and personnel staff of station are in interaction each other and they create complex system prone to harm [2]. moreover, the system is mostly part of whole transportation infrastructure connections with track and next stations, material, energetic and informatics flows, and number of served persons. it is reality that devices and fittings that are very vulnerable to damages from disasters of many kinds [3]. from the protection reasons in the practice it is necessary to identify the protected assets and to establish their criticalities taking into account all relevant disasters, i.e. the sources of hazards, and to consider possible risks. this procedure is a common part of formation of safety management systems in all systems, i.e. also in railway domain. the first step is the identification of basic metro assets that are important for safe metro operation and for human (employee and passengers) security; and the other is the determination of criticalities of these assets. 2. critical assets and critical infrastructure protection for ensuring the humans security and development, there is necessary the safe human system [3]. basic assets of human system called basic public assets are: human lives and health, property and welfare, environment, and also critical infrastructures and technologies. it means that the critical infrastructure is one of important asset of human system. therefore, its protection is specially ensured in the eu, usa, germany, canada etc. [1]. in the czech republic critical infrastructure it includes nine infrastructures [1] [4], i.e. infrastructure for: • energy (electricity; gas; heat energy; oil and petroleum products) supplies, • handling with water (drinking and industrial water; secure and manage the surface water and underground water resources; waste water system), • food supplies and agriculture operation (food production; care for food; agricultural production), • health care (pre-hospital emergency care; hospital care; the protection of public health; the production, storage and distribution of pharmaceuticals and medical devices), • transportation system (road; rail; air; water), • cyber-communication and information systems (fixed telecommunication network services; services of mobile telecommunication networks, radio communication and navigation; television and satellite 29 http://dx.doi.org/10.14311/app.2016.5.0029 http://ojs.cvut.cz/ojs/index.php/app tomáš kertis, dana procházková acta polytechnica ctu proceedings communications radio; postal and courier services; internet and data services), • the banking and financial sector (management of public finance; banking; insurance; the capital market), • emergency services (fire rescue brigade cr; units of fire protection for area coverage; police of czech republic; army of czech republic; radiation monitoring; forecasting, warning and mouthpiece for the service etc.), • public administration (state administration and self-administration; social protection and employment; the state social support and social assistance; performance of judiciary and prison system). the transportation infrastructure includes road, rail, air and water infrastructure. each of mentioned infrastructures is composed from elements and line structures. the protection of transportation infrastructure in the czech republic is only concentrated to the elements according to the crisis act (act no. 240/2000 col.). it means that metro station is key asset, and therefore, we deal with its safety. transportation system provides the transportation of persons and goods. it includes all modes of transport, which, under the coordination of the various transport systems work together and creates a logistical network. the transportation network, together with energy network can be considered as the foundation of economic prosperity of the state. the transportation network consists of separate parts that make up self-reliant closed units. each unit is a unit, the basis of which is the type of transport, accompanied by complex operational and administrative parts, which forms subsystems. the characteristic feature of technological units is their specialized ability to separate traffic operation. it is necessary to recognize that the individual parts are bound by the rules of their transport logistics, which coordinates the activities. in the czech republic the transport network consists of: rail transport system; the road transport system; the air transport system; and water transport system. transport system and its subsystems are vulnerable to natural disasters, the traffic accidents, technological accidents of facilities and buildings being in their vicinity, crimes, terroristic attacks, organizing accidents, war etc. [1] [3] [4]. figure 1 shows the processes that are sources of disasters [3]. to ensure safety of each entity, i.e. each infrastructure and their components it is necessary to manage the safety by the way as is expressed by process model in figure 2 [1] [5]. the safety needs to be formed on all levels of governance [3] [4] [5], as it is depicted in figure 3. the ways of coping with risks are shown in figure 4 [3] [4] [5]. each real system is composed from various heterogeneous subsystems, links and flows among them, so-called inter and intra-dependences. subsystems figure 1. sources of disasters in the human system [3]. figure 2. process model for formation of safety of entity under account [1] [5]. figure 3. levels of safety management [3] [4] [5]. figure 4. measures used for get over risks connected with disasters [3] [4] [5]. 30 vol. 5/2016 assets of model metro station and their criticality is possible categorized as controlled and control systems in the view of control. for improving system parameters (dependability, security, safety, etc.) the support systems are implemented, for instance a protection system reducing the risks. a subsystem can be classified as safety related, the safety related system carries out very important functions which failures or malfunctions could lead to improving the risks and event to accidents [6]. from the protection reasons on the concept of integral safety of complex technological facilities in the practice it is necessary to identify the protected assets and to establish their criticalities taking into account all relevant sources of hazards. 3. data on metro in praha and metro station model stations of metro in the praha capital are not the same; they have differences due to their locations in different site conditions and demands in the time of their building. all are threatened by disasters that threaten the praha capital, i.e. its buildings, structures and facilities, and locally also by specific phenomena that are associated with local conditions, which are connected with local vulnerabilities. therefore, for research needs we developed a model station that has characteristics that are common to all the stations of reference system taking into account the praha land-use plan [7] and the praha metro documentation [2] [8] [9] [10]. the structural parts of praha metro were constructed according to the building law (to 2006 act. no. 50/1976 col., from 2006 act no. 183/2006 col.), the engineering parts and cyber technique were harmonised with demand of building law so they might obtain permission for operation. the metro operation control is performed according to the standards [11] [12] and it satisfies requirements given in standards [13] [14] [15]. the model metro station contains all elements, links and flows that are same in all metro stations because they are determined by the requirements of mentioned building law; i.e. we follow general features and do not study the details in structure, devices, fittings of both, the technical and the cyber domains. the detail studies of individual stations reveal real vulnerabilities that cannot be open to the public from the safety reasons. such detail studies have been performed for the metro operator, and they are used for metro safety upgrade. the sources of phenomena that can damage the metro system will be discussed in the next chapter in which we use the advanced principle based on all-hazard-approach [16]. considering the principle defence-in-depth described for complex technological facilities in [17] we shall mostly concentrates to the management units that keep the sufficient level of safety at normal, abnormal and critical conditions. safety management system of metro (sms) has central control system and controlled systems in each metro model station, figure 5. it is very complex distributed system that involves controlled, control and protection subsystems, their interdependences, operation states and conditions. figure 5. control and controlled system in metro model station [6]. metro station has subsystems stationary, track and on-board. on the basis of results in [6] obtained by application of principles given in [17] we select the main metro station assets in individual main parts. technological parts and some support subsystems of controlled system include assets: (1.) energetic devices: transformer substations and distribution transformers, (2.) communication equipment: communication cables, vhf connection with trains; passengers information systems including automatic check-in; communication systems for passengers and staff; cctv, telephone, public address; clocks and fire alarms; protection systems and signalling, (3.) machinery: escalators; pump stations in nodes in stations and between them; elevators; workshops and warehouses maintenance stations, (4.) air conditioning, (5.) mobile machines and devices: rolling-stock; devices and substances for cleaning, including the containers, cleaning machines, ladders, scaffolding for cleaning the lighting systems; fire protection equipment, (6.) next important equipment: security keys and alarm buttons; equipment for fire alarm start; traction devices and lighting; track devices, the main water shut; moving stairs, plates, signal panel for machinery devices; closing equipment (shutters). praha metro control centre has denotation operation control centre occ. it has supervisory control and data acquisition (scada) system that monitors technological facilities and processes and enables their control. its variant used in the czech republic has denotation asdř-d (automatic dispatcher system for transport control). the scada has several parts that are responsible, for: transport (ugtms urban guided transport management and command / control system); energetic tasks; technologies; communication and lighting. the control system involves following assets: • train dispatcher (surveillance and control of trains); and occ nodes, station nodes, nodes for automatic route setting system – asdř-d, 31 tomáš kertis, dana procházková acta polytechnica ctu proceedings • energetic dispatcher – asdř-e, • technological dispatcher – asdř-t, • lighting system dispatcher – adrř-o, • communiation and protection systems dispatching, • fire dispatching, • depot dispatching for train services and maintenance. for a model metro station, the control system involves station that are also assets: • station nodes of control system, • station node for automatic route setting, • station nodes for connection of energetic and technological dispatching, • station nodes for central lighting control system, • station interacting with dispatchers (communication, protection, operation, fire men). in the railway practice context a protection system is primary understand as system needed to mitigation of risks connected with train operation. it is also asset. they are: • station protection systems (esa 11 m), • wayside protection systems (ažd 71, esa 11 m), • train protection systems (matra and on-board units). the metro stations also ready on flows that are also assets, it goes on: • energetic flow, • information flow, • central dispatching and station nodes of control system, • central dispatching and station communication units, • station nodes of control system and protection systems, • station nodes of control system and communication system, • protection systems and technologies (on the track and on-board)-telephone connection between central dispatching (occ) and station, • telephone connection between central dispatching and trains. controlled, control and support (protection) systems are not just electronical and mechanical devices, they can be also organization processes, other communication and information parts of the system, human resources and place which shall be taken into account in the assets identification context. in the czech practice the european directives and standards are respected for safety management of railway systems; there are taken in account: common safety methods [18] [19], en 50126-1 [13], clc/tr 50126-2 [20], en 62290 [14] [15], en 92267 [21]. railway domain is regulated by european directive 2004/49/ec [18] that establishes duties for national authorities. the mentioned directive refers also to railway interoperability and safety. according to the directive regulation 402/2013 on the common safety methods for risk assessment [19], there take into account only risks connected with human harm. reason is an effort to ensuring the safety of passengers, employees, level-crossing users, unauthorised persons on the track and to ensuring the total safety on the basis of the above categories. weakness of mentioned directive is that it does not consider other assets which affect to people lives and healthy, it means public assets, important technological parts, facilities, any processes and infrastructure functions. the directives are insufficient for critical infrastructure needs and it does not provide assets protection. moreover, mentioned directives are not apply to special railway systems like metro. railway standard en 50126 [13] defines the specification and demonstration of reliability, availability, maintainability and security (rams) for a railway system including the metro. the standard defines whole life cycle including the concept and design of the system, through operation, maintenance, reconstruction, and up to decommissioning of the system. risk management is included there, but it considers only risks affect to human, it means risk r = rate (of accidents) x degree of severity (of harm); according to clc/tr 50126-2:2007 [20] in this context harm may imply: human harm (causing injures, fatalities); and also environmental harm (damage to property, spread toxic substances, other impacts, etc.). most railway safety studies tend to concentrate on human harm. (clc/tr 50126-2:2007). security requirements to integral safety improving are not still covered. standard en 62290 [14] [15] defines requirement for urban guided transportation management and control / command system which is split into 5 degrees of automation. each degree of automation (goa) has specific demands and set of functional requirements. although a standardisation commission has been established, no standard with safety requirements has been prepared [21]. better approach is included in en 62267 [21] for fully automated guided transport (augt). the standard refer to en 50126 [13] about risk management but it imposes demand on assets protection. identified protected assets by the standard are: persons (passengers, staff, external emergency services, public); and property, shall be defined by railway authority and operator (infrastructure of whole system, trains, equipment, etc.) [22]. the european standard is intended to very short group of railway system and it does not establish methods for determining the protected assets. the characteristic pattern of model metro 32 vol. 5/2016 assets of model metro station and their criticality station is composed from data mentioned above and the facts that are summarised in [11] [12] [13] [14] [15] [18] [19] [20] [21] [22]. 4. disasters considered in research since in the czech republic there is held long-term use of building codes and high quality engineering practice there are metro stations relatively well protected against design disasters of natural and technological origin that belong to the term all-hazardapproach [16], which ensures fixed safety level in these cases. the problems are connected with further disasters that belong to the term all-hazard-approach [16] but they were been introduced into practice later. it involves the following types of disasters: • so called organisational accidents [3] [4] [5] [17] because these disaster sources have been considered since the end of 80s of last century, • having the origin in cyber domain or in interdependences of different nature in system [3] [4] [5] [17] because these disaster sources have been considered since the end 90s of last century, • damaging phenomena caused by intentional human activities as crime or terroristic attack [3] [4] [5] [17] because these disaster sources come to size after 2000. from above reasons last designated disasters were not considered in original designing the metro stations, and therefore, we will pay special attention to them. data for the research we obtained from the sources on the praha metro [2] [8] [10]. all above mentioned metro station assets are threatened by disasters. for praha capital and its facilities according to [5] [7] the following disasters are relevant: (1.) the results of the processes inside and outside the earth: flood; tornado; earthquake; liquefaction of the ground; landslide, (2.) the results of the human body, human behaviour and processes in human society: epidemics; pandemics; failure of the stability of human society (welfare disruption, criminality); assault; terrorist attack; attack using the chemical, nuclear, radiological and biological (cbrne) weapons; armed conflict; war, (3.) results of the processes and activities installed by people: industrial accident; accident during the transport or storage of hazardous substances; accident at transportation; failure of critical infrastructure; failure in economy; failure in cyber infrastructure; failure of supply chains; loss of serviceability, (4.) interaction of earth and environment to activities of people: disruption of bedrock stability due to vibration; air pollution; water pollution; rapid climate changes; migration of large groups of people, (5.) internal dependences in the human system (natural or man-made): organizational accident; failure of flows of raw materials and products; disturbance in the flow of energy; failure in the information flow. 5. method of determination of model netro assets criticality to ensure the safety of model metro station it is necessary to know the criticality of whole metro system and its components, i.e. in followed case it is the model metro station [17]. as it was said above, each metro station is threatened by disasters, i.e. its safety depends on the quality of the work with the risks [5] [17]. safety and risk are together in some way in relation, but they are not complementary variables (complementary value to the safety is the criticality [17]). the criticality is a result of the exposure to risk factors. there are analysed the different types of factors: factors of human activity, decision-making and management, factors in the nondistinguishability of reality (because there are random and knowledge uncertainties). according to [17] the criticality of technological system indicates a threshold value that means it the values are below this threshold, so the status is demanded (subcritical) and vice versa. the assessment of criticality of real system is not a trivial matter, because the elements and whole have different roles for different conditions of active, reactive, amplifying or dampening (but not additive), and among the sub systems of systems of system, which is a real entity, and each model metro station. therefore, in accordance with [17] for its determination we use the method of risk engineering the checklist [23], which for complex technological systems is given in table 1. because each system of systems (in our case metro station) is a very complex system, the behaviour of which cannot be easily predicted, so the assessments from the perspective of the different sectors are often different, and therefore, it is necessary to choose the method that conflicts does not override, but it is trying to reasonably consider them [23]. therefore, it was used the method, based on the expert opinion of five experts who knew enough the metro and its operation, have been independent and were from the areas: • safety and protection of human health at work, • protection of property, • economy of operation, • protection of passengers, • technical and functional safety of operation. the evaluation was carried out separately for each significant disaster using the value scale listed in table 2. 33 tomáš kertis, dana procházková acta polytechnica ctu proceedings i question asst note 1 1. has the technological facility to incorporate the principles of inherent safety? 2. has the control system of a technological facility (sms) set the basic control functions, alarms and the response of the operator set up so that the technological facility in normal (steady) state? 3. has management system (sms) instrumentation (built-in safety instructions) and relevant physical barriers, which at derogate from the normal state to keep technological system in a good condition, i.e. they prevent the occurrence of unwanted phenomenon? the operation is successful, when, after the occurrence of the abnormal state the technological facility will return to normal as a result of resilience or after the application of corrective measures (clean-up, repair, replacement of parts). 4. has management system (sms) for the case of loss of control, i.e. critical conditions measures for emergency response that mitigate impacts on technological facility system and ensure the capability to return to a normal state? operation of a technological object is successful, if it is a good continuity plan, which ensures that the technological facility shall ensure all the necessary tasks. 5. does management system (sms) for the case of loss of control, i.e. supercritical (beyond design, extreme) conditions the measures for: maintaining the operability of the technological system following its repair and maintenance; item and measures to ensure the protection of public assets (people, the environment and other assets) in the surroundings of technological facility? . ............................................................................................................................. ........ ..... n 1. has the technological facility to incorporate the principles of inherent safety? 2. has the control system of a technological facility (sms) set the basic control functions, alarms and the response of the operator set up so that the technological facility in normal (steady) state? 3. has management system (sms) instrumentation (built-in safety instructions) and relevant physical barriers, which at derogate from the normal state to keep technological system in a good condition, i.e. they prevent the occurrence of unwanted phenomenon? the operation is successful, when, after the occurrence of the abnormal state the technological facility will return to normal as a result of resilience or after the application of corrective measures (clean-up, repair, replacement of parts). 4. has management system (sms) for the case of loss of control, i.e. critical conditions measures for emergency response that mitigate impacts on technological facility system and ensure the capability to return to a normal state? operation of a technological object is successful, if it is a good continuity plan, which ensures that the technological facility shall ensure all the necessary tasks. 5. does management system (sms) for the case of loss of control, i.e. supercritical (beyond design, extreme) conditions the measures for: item maintaining the operability of the technological system following its repair and maintenance; item and measures to ensure the protection of public assets (people, the environment and other assets) in the surroundings of technological facility? table 1. identification of deficiencies for specific disasters in a given territory, i = 1, 2, ..., n, i.e. assessment of the criticality of the viewpoint of the application of access the defence-in-depth. 34 vol. 5/2016 assets of model metro station and their criticality criticality rate values [%] extremly high – 5 more than 95 very high – 4 70-95 high – 3 45-70 middle – 2 25-45 low – 1 5-25 negligible – 0 less than 5 table 2. value scale according to [17]. the resulting rates of criticality for each item and disaster are set as median [23] from particulars determined by individual experts. 6. criticality of assets in model metro station on the basis of the procedure described in the previous chapter it was obtained a set of values that are listed in table 3. the numbers in table 3 indicate the following rates of criticality: • 0 – given disaster does not have a direct impact on the metro station and its operation, • 1 – given disaster has low impact on the metro station and its operation, • 2 – given disaster has middle impact on the metro station and its operation, • 3 – given disaster has high impact on the metro station and its operation. from table 3, it follows that the rate of criticality is high in many cases. while it is moderate, i.e. it is not very high or extreme, it is necessary in the frame of public interest to supplement the measures in all cases where the criticality rate is high. on the basis of the data referred to such cases according to [17] [24] it is necessary to perform in all cases where the criticality rate is 3 the following: • the analysis of the ability/competence of operator of metro to handle with emergency situations in the station and in the infrastructure, • the inventory of available resources for quick and proper response to the emergency situation in a metro station in the subway, • the assessment of current vulnerability of items of metro stations and the entire metro subway, • the specification of the coordination of activities to support the continuity of subway stations and underground infrastructure, • the division of responsibilities for actions to promote the continuity of the subway stations and underground infrastructure, • the setting up the organisational instructions and features to support the renovation of subway stations and underground infrastructure, • the establishment of criteria for the selection of main features of elements of subway stations and underground infrastructure, • the determination of support processes that support the main features of subway infrastructure, • identification of key personnel, to ensure the continuity of the subway stations and underground infrastructure, • prioritizing the main functions on the basis of the criticality of time, sequence of key recovery processes and personnel capacity. 7. conclusion from the protection reasons on the concept of integral safety of complex technological facilities in the practice it is necessary to identify the protected assets and to establish their criticalities taking into account all sources of hazards. the results of performed research show that metro stations have a certain level of safety (according to [17] it holds that level (rate) of safety = 1 – criticality rate). however, it is necessary to carry out the measures referred to in the previous chapter, because there are items for which in the occurrence of large (i.e. beyond design extreme) disasters there are get to significant impacts of disasters, which have the ability to disrupt seriously the metro station and the entire infrastructure. it means that present standards and legislative do not provide sufficient range for integral safety, including the safety of important assets and systems that predispose sufficient human security and human society development. references [1] d. procházková. bezpečnost kritické infrastruktury, 2012. isbn 978-80-01-05103-0. [2] dopravní podnik v praze. dp kontakt: časopis zaměstnancøu dopravního podniku hl. m. prahy, akciové společnosti, 1999-2005. isbn 1212-6349. [3] d. procházková. strategické řízení bezpečnosti a udržitelného rozvoje území, 2011. isbn 978-80-7251-243-0. [4] d. procházková. krizové řízení pro technické obory, 2013. isbn 978-80-01-05292-1. [5] d. procházková. analýza a řízení rizik, 2011. isbn 978-80-01-04841-2. [6] t. kertis. bezpečnostní plán vybrané stanice metra, diploma thesis. praha 2015. [7] praha, hl. m. archives for strategis and land-use planning. [8] metroweb [online], 2015. [2015-03-15], http://www.metroweb.cz/. [9] kolektiv pracovníkøu metroprojektu praha a.s. publikace iv. [online], 2008. http://www.praha.eu/ public/0/a3/da/186337_4_publikace_iv._c2.pdf. [10] unicontrols a.s. case study: metro praha, 2014. 35 http://www.metroweb.cz/ http://www.praha.eu/public/0/a3/da/186337_4_publikace_iv._c2.pdf http://www.praha.eu/public/0/a3/da/186337_4_publikace_iv._c2.pdf tomáš kertis, dana procházková acta polytechnica ctu proceedings assets/disasters f lo od e ar th qu ak e l iq ue fa ct io n of gr ou nd g as re le as e e pi de m ic p an de m ic w el fa re di sr up ti on c ri m e in te nt di sr up ti on t er ro ri st ic at ta ck a tt ac k w it h c b r n e a rm ed co nfl ic t w ar in du st ri al ac ci de nt l ea ka ge of ha za rd ou s t ra ffi c ac ci de nt b la ck ou t d is ru pt io n of ec on om y human lives and healthy 1 1 1 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 1 1 1 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 1 1 1 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 property and places 1 1 1 2 0 0 0 2 3 3 3 3 3 1 3 2 2 0 1 1 1 2 0 0 0 2 3 3 3 3 3 2 3 3 2 0 2 1 1 2 0 0 0 2 3 3 3 3 3 2 3 3 3 0 energetic devices 1 1 1 2 0 0 0 2 3 3 3 3 3 1 3 2 2 0 1 1 1 2 0 0 0 2 3 3 3 3 3 2 3 3 2 0 2 1 1 2 0 0 0 2 3 3 3 3 3 2 3 3 2 0 communication equipment 2 2 2 2 0 0 0 2 3 3 3 3 3 1 3 2 3 0 2 2 2 2 0 0 0 2 3 3 3 3 3 2 3 3 2 0 2 2 2 2 0 0 0 2 3 3 3 3 3 2 3 3 2 0 machinery 1 1 1 2 0 0 0 2 3 3 3 3 3 1 3 2 2 0 2 2 2 2 0 0 0 2 3 3 3 3 3 2 3 3 2 0 3 3 3 3 0 0 0 3 3 3 3 3 3 3 3 3 2 0 air conditioning 1 1 1 2 0 0 0 2 3 3 3 3 3 2 3 2 2 0 2 2 2 2 0 0 0 2 3 3 3 3 3 2 3 3 2 0 3 3 3 3 0 0 0 3 3 3 3 3 3 3 3 3 3 0 mobile machines and devices 1 1 1 1 0 0 0 3 3 3 3 3 3 3 3 3 3 0 1 1 1 2 0 0 0 2 3 3 3 3 3 2 3 3 2 0 2 1 1 3 0 0 0 3 3 3 3 3 3 3 3 3 0 0 next important equipment 1 1 1 1 0 0 0 3 3 3 3 3 3 3 3 3 3 0 1 1 1 2 0 0 0 2 3 3 3 3 3 2 3 3 2 0 2 1 1 3 0 0 0 3 3 3 3 3 3 3 3 3 0 0 station nodes of control systems of metro system 1 1 1 0 0 0 1 1 1 1 1 1 1 1 3 1 1 0 1 1 1 0 0 0 2 1 1 2 1 1 2 1 3 1 1 0 3 2 2 0 0 0 3 2 2 3 3 2 2 2 3 2 2 0 protection systems 1 1 1 0 0 0 1 1 1 1 1 1 1 1 3 1 1 0 1 1 1 0 0 0 2 1 1 2 1 1 2 2 3 1 1 0 3 2 2 0 0 0 3 2 2 3 3 2 3 3 3 2 2 0 linkages and flows 1 1 1 0 0 0 1 1 1 3 3 3 3 3 3 3 3 3 1 1 1 0 0 0 2 1 1 3 3 3 3 3 3 3 3 3 3 2 2 0 0 0 3 2 2 3 3 3 3 3 3 3 3 3 table 3. criticality of assets in metro station. [11] čr. čsn en 13816 (269389) doprava – logistika a služby – veřejná přeprava osob – definice jakosti služby, cíle a měření, 2003. [12] čr. čsn en iso 9001:2009 (01 0321) systémy managementu kvality – požadavky, 2009. [13] čr. čsn en 50126-1 (333502) drážní zařízení – stanovení a prokázání bezporuchovosti, pohotovosti, udržovatelnosti a bezpečnosti (rams): část 1: základní požadavky a generický proces, 2001. [14] čr. čsn en 62290-1:2006 (33 3530) drážní zavrízení – systémy řízení městské dopravy s vyhrazenou vodicí dráhou: část 1: systémové principy a základní pojmy, 2007. [15] čr. čsn en 62290-2:2011 (33 3530) drážní zařízení – systémy řízení městské dopravy s vyhrazenou vodicí dráhou: část 2: specifikace funkčních požadavkøu, 2009. [16] fema. guide for all-hazard emergency operations planning [online], 1996. http://www.fema.gov/pdf/plan/slg101.pdf. [17] d. procházková. bezpečnost složitých technologických systémøu, 2015. isbn 978-8001-05771-1. [18] eu. directive 2002/49/ec of the european parliament and of the council of 25 june 2002 relating to the assessment and management of environmental noise – declaration by the commission in the conciliation committee on the directive relating to the assessment and management of environmental noise, 2002. 36 http://www.fema.gov/pdf/plan/ slg101.pdf vol. 5/2016 assets of model metro station and their criticality [19] eu. regulation 402/2013 on the csm for risk assessment and repealing regulation 352/2009, 2002. [20] celenec. clc/tr 50126-2:2007 railway applications. the specification and demonstration of reliability, availability, maintainability and safety (rams). guide to the application of en 50126-1 for safety, 2013. [21] iec. project iec 62290-3 ed. 1.0: railway applications – urban guided transport management and command/control systems – part 3: system requirements specifications, 2015. [22] čr. čsn en 62267 (33 3532) drážní zařízení – automatizovaná městská doprava s vyhrazenou vodící dráhou (augt) – bezpečnostní požadavky, 2009. [23] d. procházková. metody, nástroje a techniky pro rizikové inženýrství, 2011. isbn 97880-01-04842-9. [24] d. procházková. základy řízení bezpečnosti kritické infrastruktury, 2013. isbn 978-80-01-05245-7. 37 acta polytechnica ctu proceedings 5:29–37, 2016 1 introduction 2 critical assets and critical infrastructure protection 3 data on metro in praha and metro station model 4 disasters considered in research 5 method of determination of model netro assets criticality 6 criticality of assets in model metro station 7 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0162 acta polytechnica ctu proceedings 13:162–165, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app wang tiles local tilings created by merging edge-compatible finite element meshes lukáš zrůbek∗, anna kučerová, martin doškář department of mechanics, faculty of civil engineering, czech technical university in prague, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: lukas.zrubek@fsv.cvut.cz abstract. in this contribution, we present the concept of wang tiles as a surrogate of the periodic unit cell method (puc) for modelling of materials with heterogeneous microstructures and for synthesis of micro-mechanical fields. the concept is based on a set of specifically designed cells that compresses the stochastic microstructure into a small set of statistical volume elements – tiles. tiles are placed side by side according to matching edges like in a game of domino. opposite to the repeating pattern of puc the wang tiles method with the stochastic tiling algorithm preserves the randomness for reconstructed microstructures. the same process is applied to obtain the micro-mechanical response of domains where the evaluation as one piece would be time consuming. therefore the micro-mechanical quantities are evaluated only on tiles (with surrounding layers of tiles of each addressed tile included into the evaluation) and then synthesized to the micro-mechanical field of whole domain. keywords: heterogeneous microstructures, wang tiles, micro-mechanical fields, local tiling, finite element method. 1. introduction the heavily increasing pressure to the materials utmost performance leads to necessity of excellent understanding of characteristic mechanical processes taking place at a microstructural level of materials. therefore the numerical material models needs to incorporate also the micro-scale level. however, even though the materials may look like homogeneous from macro-scale view, the majority of materials used in engineering applications are heterogeneous with a stochastic microstructure layout (fig.1). for modelling of heterogeneous materials many methods exists, but the most widely used one is the numerical homogenization method [1]. heterogeneous materials are modelled using the concept of periodic unit cell (puc) or statistically equivalent periodic unit cell (sepuc). the puc approach is used for material with periodic patterns, where the modelled microstructure is produced by duplication of a representative part of the original microstructure in cardinal directions (see fig.2). on the contrary the stochastic microstructures are modelled using the concept of sepuc. in this case the unit cell representing the original microstructure holds the same properties given by statistical descriptors as the source one [2]. nevertheless, this approach uses only a single cell to describe the whole material and the stochastic heterogeneous material is transformed into unwanted periodic pattern of the same cells. and to be able to study micro-mechanical quantities (e.g. stresses, strains or displacements) at micro-scale level we need to preserve not only its properties but also the stochastic layout. (a) (b) figure 1. the appearance of the structure at different scales, a) homogeneous at macro-scale, b) heterogeneous at micro-scale. (a) (b) (c) figure 2. modelling of periodic microstructure, a) periodic microstructure, b) puc, c) reconstructed microstructure. 162 http://dx.doi.org/10.14311/app.2017.13.0162 http://ojs.cvut.cz/ojs/index.php/app vol. 13/2017 local tilings from merged edge-compatible fe meshes (a) (b) 1 3 5 7 8 figure 3. principle of wang tiles method, a) domino game, b) wang tiles. preserving the randomness of the recreated material cannot be achieved with single cell. hence the method utilizing the set of representative cells must be used. to be able to place these cells side by side, they must be specifically designed so they are mutually compatible on the edges. if this criterion is met the infinitely large material microstructure can be created with the use of simple algorithm. the resulting microstructure preserves stochastic distribution and the statistical description of the source media. the above described method is called method of wang tiles. 2. wang tiles method the basic principle of wang tiles method was introduced by mr. hao wang [3, 4] in the early sixties of the 20th century. the concept is very similar to the domino game. the objective is to construct a straight or zigzag line series of game pieces with a matching number of dots on the adjacent sides of touching stones (fig. 3a). the wang tiles method works nearly the same with only few differences. instead of double coded rectangular domino pieces, the tiles are squares with four edge codes, colours in this case (fig. 2b). and contrary of creating a linear series of game pieces the tiles are placed into four cardinal directions (north – n, south – s, east – e, west – w) and the two-dimensional area covered with tiles is produced. the tiles cannot be rotated or mirrored through the process of laying tiles as any of these two actions would create a new type of tile. the tiles are placed sideby-side according to matching edge codes until the desired size of area is completely created. the result is called tiling and if there are no missing tiles or unequal edge codes, the tiling is valid. 2.1. cshd algorithm except the condition that the tiling must be valid we also need it to be stochastic. that can be achieved by utilization of cohen-shade-hiller-deussen (cshd) tiling algorithm [5]. if the creation of tiling is performed row by row, the most difficult position to place a new tile would be the corner position like in figure 4. in this picture is shown the corner position where the north (n) and western (w) edges of new tiles must n w 1 3 4 8 2 6 2or ? figure 4. tiling process and valid tiles for the corner position. 1 2 3 4 5 6 7 8 figure 5. smallest wang tiles set called w8/2-2. correspond with already placed tiles. among other rules of the cshd algorithm, the most important one is that it there must be at least two valid tiles (like tiles number 8 or 2 in the fig. 4) for this kind of corner position. then the new tile is randomly selected from these valid tiles and placed in the tiling. this creates another corner position and the process is repeated until the tiling of required size is created. the smallest set that complies the requirements of cshd algorithm and has at least two distinct vertical edge codes and two distinct horizontal edge codes is set consisting of 8 tiles. we call this set w8/2-2 (fig. 5) and we use it as a representative set in all further explanations in the following paper. 2.2. design of tiles when designing the set of tiles the process must ensure that the resulting tiles will satisfy the requirement of the cshd algorithm and of course the tiles must hold the properties of the original microstructure. couple of methods to design tiles can be used and one of them is based on multicriterial optimization of tile morphology according to statistical descriptors. the goal is to create tiles such as the recreated microstructure share the same statistical properties as the original microstructure. however this process is extremely time and computationally consuming (days to weeks). moreover as presented in the paper [6] with increasing pressure on optimizations all the tiles tend to become identical and the reconstructed microstructure became visibly periodical. another method called automatic tile design [5] creates each tile by stapling four small samples (fig. 5) cut out of the original microstructure in different combination. in addition is almost instantaneous if compared with first method. completely different approach based on particles dynamics, movement and collisions inside of the interconnected tiles bounding boxes are using methods described in [7, 8]. these method can be used only for microstructures with distinguishable particles that can be described by simple geometrical shapes. 163 l. zrůbek, a. kučerová, m. doškář acta polytechnica ctu proceedings 1 3 2 4 1 2 3 4 a) b) c) figure 6. automatic tile design [5], a) cut-outs for stapling, b) stapling, c) new tile. 3. micro-mechanical quantities the wang tiles method can be used in many different fields like modelling of quasi-crystals and generating naturally looking textures for computer graphic [5]. or as a substitution for the unit cell concepts (puc and sepuc) in material modelling. or to obtain the micro-mechanical response for enrichment functions in numerous generalized finite element methods, e.g. [9, 10]. in our work, the following (numerical homogenization like) decomposition of a sought macroscopic field u(x) is assumed u(x) = ue (x) + u∗(x) (1) where u∗(x) is the microscopic fluctuation field synthesized by means of wang tiles method and ue (x) is the macroscopic part solved by finite element method (fem) with a sparse grid. 3.1. simple synthesis the whole microscopic domain can of course be discretized by very fine fem mesh and evaluated to obtain demanded micro-mechanical fields like stresses, strains and displacements. but according to the fineness of the mesh and size of the domain, the evaluation time is rapidly increasing. the computational requirements can be diminished with the wang tiles method as the process of microstructure reconstruction can be applied to the micro-mechanical field. the micro-scale quantities are evaluated on individual tiles and then synthesized back according to the same tiling as the microstructure. however, as presented in the paper [11] the non-local character of mechanical quantities is causing discontinuities on tile edges and therefore the underlying grid of tiles is recognizable. that is because of the mechanical response of each tile is affected by a different combination of surrounding tiles (fig. 7). 3.2. finite element meshes furthermore, to be able to use the finite element method (fem) as a solver brings another difficulty. because the domain is represented by the tiles which are mutually compatible on the edges, the finite element meshes created on the individual tiles has to fulfil the exactly same compatibility. example of artificially created edge-compatible finite element meshes on tiles are shows in the figure 8. 2 1 6 62 6 3 4 8 3 6 4 5 8 7 1 8 5 4 511 1 1 1 1 12 2 2 2 2 2 2 2 2 2224 4 4 4 4 4 4 8 8 8 88 8 8 3 3 33 3 3 5 5 5 5 55 5 7 7 7 7 7 7 77 7 6 6 6 6 1 6 m=3 figure 7. single layer of surrounding tiles (dashed squares) for different tiles with no. 1. 1 2 3 4 5 6 7 8 figure 8. mutually edge-compatible finite element meshes on tiles. 3.3. local tiling our attempt to solve this problem is to include these surrounding tiles into the evaluation of the mechanical response for each tile [12]. therefore for every tile ti (grey square in the in figure 9) from the macro tiling we create so called local tiling. this local tiling is represented by the center tile ti and the arbitrary number of surrounding layers of tiles taken from the macro tiling. then the local local fe mesh is synthesized (see section 3.4) and the micro-mechanical quantity is evaluated. because we need only the results for the centre tile ti we crop the results of all surrounding tiles and save the results under label t285416148. this procedure is repeated for every tile ti in macrotiling and because these processes are not dependent, they can be solved using parallel computation. after results of all tiles are obtained, the micro-mechanical field is synthesized. 3.4. fe meshes merging tool as mentioned in section 3.2 the fe meshes on tiles must be mutually edge-compatible so they can be merged together into single fe mesh (e.g. local tiling fe mesh). to create the edge-compatible meshes 164 vol. 13/2017 local tilings from merged edge-compatible fe meshes evaluated mechanical response 2 8 5 4 81 4 6 1 2 8 4 4 8 5 6 1 7 1 8 5 4 51 1 1 1 2 2 2 22 4 4 4 4 8 88 3 3 5 5 7 7 7 7 77 6 6 6 1 812 2 2 57 2 2 4 5 7 6 11 2854 61481 figure 9. local tiling process: macro tiling ⇒ local tiling ⇒ local fe mesh ⇒ evaluation ⇒ cropping results ⇒ results for tile ti. figure 10. real adaptive mesh of tile. the software called t3d is used. next the output from t3d must be loaded, sorted, renumbered, etc. for these purposes we are working on robust tool implemented in c++ language. in addition to above mentioned operations, the tool can already merge arbitrary number of tile meshes (without duplicate nodes or gaps in node numbering). any created mesh can be saved to a file or loaded from file (once saved) and especially visualized using the vtk visualization tool-kit as an unstructured grid. the tool is capable to process any finite element mesh from simple (fig. 8) to quite complex like on fig. 10. the entire current work is devoted to implement the process described on figure 9 and on the effective storing and working with big volume of data results. 3.5. future work once the implementation of the local tiling process (see 3.3) is completed and we will be able to quickly generate results we are planning to implement some statistical descriptors, automatized call of t3d software or even tile design methods. our main goal is to run sensitive analysis and determine the influence of characteristic microstructural lengths, tile size and number of necessary included layers in local tilings on the overall error of synthesized micro-mechanical fields. acknowledgements the authors gratefully acknowledge the financial support from the grant agency of the czech technical university in prague, the grant no. sgs17/042/ohk1/1t/11 (numerical methods for modelling uncertainties in civil engineering) (l. zrůbek, a. kučerová and m. doškář). references [1] f. fritzen. microstructural modeling and computational homogenization of the physically linear and nonlinear constitutive behavior of micro-heterogeneous materials, vol. 1. kit scientific publishing, 2011. [2] h. lee, m. brandyberry, a. tudor, k. matouš. three-dimensional reconstruction of statistically optimal unit cells of polydisperse particulate composites from microtomography. phys rev e 80:061301, 2009. doi:10.1103/physreve.80.061301. [3] h. wang. proving theorems by pattern recognition–ii. bell system technical journal 40(1):1–41, 1961. doi:10.1002/j.1538-7305.1961.tb03975.x. [4] h. wang. games, logic and computers. scientific american 213(5):98–106, 1965. [5] m. f. cohen, j. shade, s. hiller, o. deussen. wang tiles for image and texture generation. acm trans graph 22(3):287–294, 2003. doi:10.1145/882262.882265. [6] j. novák, a. kučerová, j. zeman. compressing random microstructures via stochastic wang tilings. phys rev e 86:040–104, 2012. doi:10.1103/physreve.86.040104. [7] d. šedlbauer, m. lepš. dynamic packing of stochastic wang tiling for heterogeneous material modelling. in proceedings of the 20th international conference on computer methods in mechanics, pp. 447–448. 2013. [8] j. stránský, m. doškář. stochastic wang tiles generation using the discrete element method. in engineering mechanics 2015, vol. 821 of applied mechanics and materials, pp. 392–398. trans tech publications, 2016. doi:10.4028/www.scientific.net/amm.821.392. [9] j. m. melenk, i. babuška. the partition of unity finite element method: basic theory and applications. computer methods in applied mechanics and engineering 139(1):289–314, 1996. doi:10.1016/s0045-7825(96)01087-0. [10] t. de freitas j.a. formulation of elastostatic hybrid-trefftz stress elements. computer methods in applied mechanics and engineering 153(1):127–151, 1998. doi:10.1016/s0045-7825(97)00042-x. [11] j. novák, a. kučerová, j. zeman. microstructural enrichment functions based on stochastic wang tilings. modelling and simulation in materials science and engineering 21(2):025–014, 2013. doi:10.1088/0965-0393/21/2/025014. [12] l. zrůbek, a. kučerová, j. novák. synthesized fields fluctuations by means of wang tiles and local tilings. in modern methods of experimental and computational investigations in area of construction ii, vol. 1144 of advanced materials research, pp. 178–183. trans tech publications, 2017. doi:10.4028/www.scientific.net/amr.1144.178. 165 http://dx.doi.org/10.1103/physreve.80.061301 http://dx.doi.org/10.1002/j.1538-7305.1961.tb03975.x http://dx.doi.org/10.1145/882262.882265 http://dx.doi.org/10.1103/physreve.86.040104 http://dx.doi.org/10.4028/www.scientific.net/amm.821.392 http://dx.doi.org/10.1016/s0045-7825(96)01087-0 http://dx.doi.org/10.1016/s0045-7825(97)00042-x http://dx.doi.org/10.1088/0965-0393/21/2/025014 http://dx.doi.org/10.4028/www.scientific.net/amr.1144.178 acta polytechnica ctu proceedings 13:161–164, 2017 1 introduction 2 wang tiles method 2.1 cshd algorithm 2.2 design of tiles 3 micro-mechanical quantities 3.1 simple synthesis 3.2 finite element meshes 3.3 local tiling 3.4 fe meshes merging tool 3.5 future work acknowledgements references tm-2018.pdf acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0006 acta polytechnica ctu proceedings 11:6–11, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app development of the municipal noise-absorbing screen and test section construction technology ondřej bret∗, petra čížková czech technical university in prague, faculty of civil engineering ∗ corresponding author: ondrej.bret@fsv.cvut.cz abstract. the article presents the development and construction technology of the municipal noise-absorbing screen (mnas) – in czech "městská protihluková clona" (mphc). the municipal noise-absorbing screen is a new acoustical element developed in cooperation of the department of railway structures, faculty of civil engineering, ctu in prague and the montstav cz s.r.o. company. it is characterised by very small dimensions – its height is only ca 0.3 metre above the top of rail. the element serves to attenuate the noise induced by rail traffic arising at the wheel-rail contact. the municipal noise-absorbing screen consists of individual, mutually interlocked units, placed one after another, made of bound recycled rubber. the article describes the major limiting conditions for the placement, shape and execution of the municipal noise-absorbing screen. it also describes the construction of the test section situated on a tram track in prague and sums up the experience obtained after the first year of operation. keywords: municipal noise-absorbing screen, noise pollution, tram traffic, tram cars, recycling, rubber granulate. 1. introduction the present-day requirements for admissible noise levels induced by both road and railway traffic in builtup residential areas have been increasingly growing. this issue is a subject of long-term discussions not only among professionals, but more and more frequently also within the public media space. the problem arising during the approval of reconstructed tram tracks in 2017 can serve as an example [1]. traffic induced noise can be limited by active or passive noise reduction measures. active noise reduction measures on the railway track in construction terms include, above all, a high-quality and well-maintained roadway, appropriate track construction design, rail fastening, application of rail absorbers and elimination of rail corrugation [2]. active noise reduction measures also include measures performed on vehicles. these are, above all, the wheel profile maintenance, bogie-mounted plates and other modifications of bogies. even the application of suitable combinations of active measures, however, comes up against the ceiling of its potential and passive measures are necessary. passive noise reduction measures include, for example, classic noise barriers, applied not only in rail transport, but also in road transport. the latest efforts are aimed at developing smaller-size elements, placed closed to the source of noise taking the form of various low noise-absorbing screens [3]. low noise-absorbing screens are elements or building units placed along tram tracks, railway tracks or other rail guided transport systems whose application results in the reduction of undesirable effects of noise arising during the passage of rail traffic [2]. low noiseabsorbing screens are placed as close as possible to the structure gauge set by a technical standard. thanks to their dimensions, they can be installed much closer to the source of noise as compared to classic noise barriers being thus able to attenuate the dominant noise arising at the wheel-rail contact (rolling noise) and the bogie noise [4, 5]. depending on the material from which they are produced, they can have absorption or reflection properties. 2. municipal noise-absorbing screen the municipal noise-absorbing screen (hereinafter also referred to as "mnas") is a noise-absorbing element developed in cooperation of the department of railway structures, faculty of civil engineering, ctu in prague and the montstav cz s.r.o. company, engaged in the recycling and production of rubber elements. the operational verification runs in cooperation with the prague public transit company (pptc). mnas serves for the attenuation of the noise induced by rail traffic, arising at the wheel-rail contact. the characteristic height of the screen is ca 30 cm above the top of rail (tr). the screen is placed in the immediate vicinity of the vehicle gauge and thus it effectively reduces the arising noise. the municipal noise-absorbing screen is characterised by the material used – a composite based on recycled rubber granulate, bound by a polyurethane binder. the rubber granulate is a secondary raw material that can be used in railway construction in different applications [6], either in a bound [7, 8] or unbound state [9]. unlike other similar elements, mnas is easily demountable, which is a great advantage in the case of e.g. a nec6 http://dx.doi.org/10.14311/app.2017.11.0006 http://ojs.cvut.cz/ojs/index.php/app vol. 11/2017 development of the municipal noise-absorbing screen essary intervention of the integrated rescue system for which concrete noise barriers create a significant obstacle to a fast and smooth intervention [10]. the first-generation mnas, which is the subject of this article, is presently being verified under operating conditions and its durability is monitored – due to favourable thermal and technical properties of the rubber granulate [9] mnas is presumed to resist winter conditions and degradation quite well [11]. it is made up of units (segments) manufactured of bound recycled rubber with dimensions of 250 x 600 x 750 mm (width x height x length), which are partially embedded below the ballast bed so that the height of the screen is ca 30 cm above the top of rail depending on the execution. 3. limiting conditions for mnas placement, shape and execution there were many limiting factors for the placement, shape and execution of the municipal noise-absorbing screen emerging already in its design phase. the major limitations are the structure gauge, or the tramway vehicle gauge respectively. the design of the mnas shape and placement was initially based on the vehicle gauge under čsn 28 0318 – "structure gauges of tramway tracks" [12]. according to this standard, the vehicle gauge width in a straight track section is 1300 mm from the rail axis. the reference vehicle considered in the above standard is a three-section tram of the kt8d5n type with a vehicle width of 2.5 m and a length of 30.3 m. in prague, this type is represented by a bi-directional tram of the kt8d5.rn2p type, which originated by the reconstruction of the kt8d5 tram [13, 14]. due to the fact that the vehicle gauge is very significantly extended in directional curves of small radii [12], the installation of mnas is feasible mainly in straight sections and curves with radii from ca 200 m. in straight sections and curves with large radii, it must be placed respecting the vehicle gauge pursuant to čsn 28 0318 [12], or other limitations specific for the respective tram operation. other parameters for the mnas placement on a test section within the prague public transit company (pptc) network included potential passage of the psp-01 bogie snow plough (figure 1). the psp01 snow plough was developed to meet the growing need for clearing snow from the rail space for the use of new low-floor vehicles. due to the low-floor part, these vehicles have a greater width in the lower parts of the body casing than older (high-floor) trams, being, in some cases,equipped with other components in the part close to the top of rail which may cause problems in contact with water (snow). to allow smooth snow removal even in directional curves, the share of the psp-01 bogie plough has a greater width than standard operated trams. its total width is 2.6 m – i.e. 1.3 metres from the rail axis. based on the need to preserve the possibility of the plough passage, this parameter must also be taken into account in figure 1. psp-01 snow plough [15]. prague and the screen must be placed with some reserve slightly further away from the rail axis than would be necessitated by the tram vehicle gauge. based on consultations with pptc representatives over designs taking into account the above limitations, a question arose concerning a potential passenger escape exit from the tram in case of emergencies (accidents, vehicle failures, etc.,...) with respect to the screen placement. while older trams and their retrofits usually have four-section folding doors (opening in the inside direction) and modern trams are, as a rule, fitted with ejecting doors (opening in the outside direction, but only to a distance of several cm from the body casing), the possibility of opening tilting doors in the t6a5 type trams appeared as potentially problematic interfering with the screen placement. as seen in figure 2, while opening the door the door leaves move along a circular trajectory getting to a distance of up to 360 mm from the body casing (1610 mm from the rail axis). therefore, when placing mnas to the right in the riding direction, the opening of the t6a5 tram door must be allowed, which can be achieved in two ways: 1) move the municipal noise-absorbing screen further way from the rail axis, to a distance exceeding 1610 mm – which is a solution totally inappropriate it terms of the acoustic effectiveness of the mnas element, or 2) select such a height of the screen above the top of rail allowing the tilting of the door – due to the height of the door at the maximum admissible wheel wear this level is at a height of 28 cm above tr – which is a solution feasible without problems. in placing the screen in the test section, this issue did not need to be solved as mnas is placed to the left of the rail axis there, i.e. on the side without a door where it does not affect a potential evacuation of the passengers. it should be added that t6a5 trams are gradually discarded from traffic operation in prague and replaced with new vehicles in which the door opening is solved by the above mentioned ejection of leaves, so that this limitation in placing mnas on the door side will no longer apply in the future. based on the above conditions and limitations, a 7 ondřej bret, petra čížková acta polytechnica ctu proceedings figure 2. t6a5 tram – cross section, including the tilting door movement – plotting in a type drawing [16]. figure 3. mounting of mnas prototypes – vandal resistance test. design of the mnas placement in characteristic sections was developed and verified by the production and mounting of three prototype segments. 3.1. verification of prototypes the trial placement of the prototypes was carried out on 23. 3. 2016 on a tram track close to the pptc central workshops, in a section between ústřední dílny dopravního podniku – na homoli stops. both the team of the department of railway structures fce ctu and the representatives of pptc, tt (tram track) administration section and dct (tram traffic route) special structures participated in the test. apart from the construction / installation test, the verification included the compatibility and interaction of individual units and their vandal-resistance was tested – resistance to the displacement of individual segments of the screen (figure 3). based on local investigation, the final design of the placement and shape was developed and the connection of mnas units was optimized. 3.2. selection of a section for trial installation the test section for the placement and verification of the acoustic properties of the municipal noiseabsorbing screen (mnas) was selected in cooperation with the prague public transit company (pptc) again. tracks with an open track structure were considered (even though mnas can also be installed in a section with vegetation cover) of which the final choice fell on a tram track in prague 4, direction sídliště modřany, the section pobřežní cesta – přístaviště in the prague district of braník. the double track tram line in the respective section is led parallel to modřanská street, on a separate body with an open ballast bed. grooved rails of the nt1 shape on tb93 concrete sleepers with w14 elastic baseplateless fastening are used there. the rail is executed as continuous welded rail. the screen was installed on the centre-bound track, to the left of the riding direction (i.e. in the inter-track space). the mnas position was selected with regard to the subsequent technical noise level measurement aimed at identifying the acoustic effectiveness of the designed noise reduction measure. the sound level meter microphone was located at a distance of 7.5 metres from the monitored track, on the side more distant from the road to eliminate the effect of passing car traffic. the characteristic section at the point of the noise-absorbing screen placement is shown in figure 4. the selected distance of the screen face from the rail axis was 1420 mm, and its height was 300 mm above the top of rail. 3.3. test section construction the test section was established on 9. 4. 2016 in coordination with a one-day tram operation lockout in the dvorce – sídliště modřany section, planned for the overhaul and maintenance of traction wires. in the morning before the lockout, accurate measurement and marking of the site for the placement and installation of the municipal noise-absorbing screen plus other preparatory works had been performed. immediately after the start of the lockout, the mhs two-way excavator made a groove in the ballast bed for the installation of mnas units (figure 5). successive work was done manually, without machinery. the site construction camp and vehicle parking were established near pobřežní cesta stop, by the exit of the link road from barrandov bridge onto modřanská street, ca 200 m from the established test section. from there, the mnas segments, tools and other material were transported by a track trolley. the width of the previously excavated groove necessary for mounting the mnas segments and the depth related to the top of rail (tr) were checked and further extended as necessary. before the installation of units, the gravel bearing surface was compacted with a manual petrol powered vibrating plate (figure 6). individual mnas units were gradually mounted onto the compacted surface, proceeding step by step from the start of the section towards přístaviště stop. the position of installed units in relation to the rail was continuously checked – their distance from the rail axis (structure gauge) and their height above tr (figure 7). mounted units were gradually backfilled 8 vol. 11/2017 development of the municipal noise-absorbing screen figure 4. characteristic section at the point of the noise-absorbing screen placement. figure 5. groove excavation for mounting mnas using the mhs two-way excavator. figure 6. compaction of groove surface prior to the screen placement. with gravel, which was compacted, both manually and using machinery. individual units were mutually interlocked by wooden laths placed on the upper side of the screen (figure 8) – this ensures the interaction of the whole screen (necessary e.g. to resist vandalism) and also secures its permanent position. figure 7. checking the screen distance and height from the rail axis and tr. an atypically chamfered unit preventing the crash of tram units into the front of the screen in case of emergency (e.g. displacement of screen units due to vandalism) was mounted onto the front of the municipal noise-absorbing screen in the riding direction (i.e. in the direction from pobřežní cesta stop) (figures 89). after the final check of the position of all units, final modifications of the ballast bed profile were made and the gravel around the installed mnas was compacted with a manual petrol powered vibrating plate. a total of 32 mnas units were installed, the test section being 24 metres in length. 3.4. screen modification following noise measurement monitoring the state after the mnas construction, the screen was modified to improve the acoustic effectiveness of the noise reduction measure used. this modification consisted in adding boards of recycled rubber, 25 mm in thickness, to the upper screen wall, which created an overlap, 9 ondřej bret, petra čížková acta polytechnica ctu proceedings figure 8. mnas seen from the start of the section – chamfered first unit is evident in the front. figure 9. detail of modified mnas execution. 70 mm in length, at the screen face to further reduce the arising noise. thanks to the board overlap, the screen face was shifted towards the track, from the original distance of 1.42 m from the rail axis to 1.35 m after the modification. the modification is evident from figure 9. the screen was modified under traffic operation on 26.4.2016. the modification was performed using boards of recycled rubber, pre-cut to the necessary dimensions. the boards were fastened to individual mnas segments with 0.5 m spacing by means of dowels into polystyrene – this solution proved to be the most convenient based on tests. after the installation of additional boards providing the overlap of the screen towards the track, the measurement of noise was carried out to verify the effectiveness of modified mnas. the solved section with completed mnas, including the modification, is presented in figure 10. 4. conclusion the designed and implemented noise reduction measure proved to be an effective alternative for the attenuation of noise induced by tram traffic. the measured data served to establish the average energy value for all measured passages of tram units. the resulting reduction due to the municipal noise-absorbing screen (mnas) application after its modification identified figure 10. municipal noise-absorbing screen after modification. was 3.9 db. some types of trams (in relation to the body casing shape at the bogie level) even reached higher noise reduction levels. more detailed information about the mnas effectiveness will be published in the professional peer-reviewed journal, akustika, in autumn 2017 (www.akustikad.com). the test section is currently subjected to continuous monitoring and different parameters are surveyed on it on a long-term basis. after the first year of operation and the first winter season, the municipal noise-absorbing screen does not show any signs of mechanical damage, degradation or other defects. in 2016, the industrial property protection agency awarded the municipal noise-absorbing screen the industrial and utility model certificate. in the following phases of research and development, further monitoring of the first-generation mnas test section, described in this article, and the construction and verification of the second-generation mnas test section or sections is presumed where structural improvements, designed on the basis of experience obtained from the construction and monitoring of the first-generation test section, will be applied. references [1] k. frouzová. praha musí řešit hlučné tramvaje na opravených tratích. hrozí omezení nebo zrušení desítky spojů / prague must solve noisy trams on repaired tracks. there is a threat of limitation or cancellation of ten lines, in: www.ihned.cz: hospodářské noviny [online]. [cit. 2017-03-10] http://domaci.ihned.cz/c1-65569370-praha-musi-resithlucne-tramvaje-na-opravenych-tratich-hrozi-omezeninebo-zruseni-desitky-spoju. [2] p. čížková. the acoustic effectiveness of low height noise barrier. in: kropp, w, ed. proceedings of the 45th international congress on noise control engineering. inter-noise 2016, hamburg, 2016-08-21/2016-08-24. berlin: german acoustical society (dega), 2016. isbn 978-3-939296-11-9. [3] j. smutný, l. pazdera, v. tomandl. assessment of the effectiveness od low noise barrier brens barrier in tetčice village. in: eisereich, j.e., ed. rychlost s tichostí. rychlost s tichostí 2015, plzeň. praha: powerprint, 2015. isbn 978-80-87994-28-3. 10 vol. 11/2017 development of the municipal noise-absorbing screen [4] p. čížková, l. lomoz. the effectiveness of a low height noise barrier. in: 14th geoconference on ecology, economics, education and legislation. 14th international multidisciplinary scientific geoconference sgem 2014, albena, 2014-06-17/2014-06-26. sofia: stef92 technology ltd., 2014. issn 1314-2704, isbn 978-619-7105-18-6, doi 10.5593/sgem2014/b52/s20.094. [5] p. čížková, l. lomoz. comparison of efficacy of low and hight acoustic screens. akustika, 2014. issn 1801-9064. [6] m. lidmila. nové perspektivy využití recyklovaných materiálů v železničním stavitelství / new perspectives of using recycled materials in railway construction, in: udržitelná výstavba. čvut v praze, 2011. isbn 978-80-01-04961-7. [7] l. horníček, h. krejčíková, m. lidmila, et al. antivibrační rohože v železničním stavitelství v čr / antivibration mats in railway construction in the czech republic. české vysoké učení technické v praze, fakulta stavební, 2011. isbn 978-80-01-04963-1. [8] j. makuch. maty wibroizolacyjne we wrocławskich torach tramvajowych. in: xiv konferencja naukowo-techniczna "drogi kolejowe 2007" [online], 2007. [cit. 2017-03-7] http://www.zits.pwr.wroc.pl/makuch/art38.pdf. [9] m. lidmila, v. lojda, o. bret. laboratory models of a railway track with the subgrade layer based on rubber granulate in: advances and trends in engineering sciences and technologies ii: proceedings of the 2nd international conference on engineering sciences and technologies, 29 june – 1 july 2016, high tatras mountains, tatranské matliare, slovak republic. boca raton: crc press, 2016. isbn 9781138032248. [10] protihlukové stěny hluku nezabrání a jsou bezpečnostní komplikací / noise-absorbing walls do not prevent noise being an obstacle to safety, želpage [online]. [cit. 2017-03-10] http://www.zelpage.cz/zpravy/8188?lang=cs. [11] o. bret. determination of rubber granulate frost protection. in: juniorstav 2017, 19. odborná konference doktorského studia. brno: vysoké učení technické v brně, fakulta stavební, 2017. isbn 978-80-214-5473-6. [12] čsn 28 0318. průjezdné průřezy tramvajových tratí a obrysy pro vozidla provozovaná na tramvajových dráhách / structure gauges and vehicle gauges for vehicles operated on tram tracks. praha: český normalizační institut, 2015. [13] katalogový list tramvajového vozu kt8d5. skd trade, a.s. / catalogue sheet of t6a5. tram car of skd trade a.s. [online]. [cit. 2017-03-18] http://www.skd.cz/tramvaje/kt8d5.htm. [14] r. mara. pražské tramvaje. praha: pavel malkus dopravní vydavatelství, 9/2012. isbn 978-80-87047-28-6. [15] archives of the prague public transit company. [16] katalogový list tramvajového vozu t6a5. skd trade a.s. / catalogue sheet of t6a5. tram car of skd trade a.s. [online]. [cit. 2017-03-18] http://www.skd.cz/tramvaje/t6a5.htm. 11 acta polytechnica ctu proceedings 11:6–11, 2017 1 introduction 2 municipal noise-absorbing screen 3 limiting conditions for mnas placement, shape and execution 3.1 verification of prototypes 3.2 selection of a section for trial installation 3.3 test section construction 3.4 screen modification 4 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2018.18.0028 acta polytechnica ctu proceedings 18:28–31, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app basic biomechanical characterization of polyurethane based artificial cancellous structures jan šleichrta, b, ∗, daniel kytýřa, kateřina pithartováa, b, sascha senckc, david fürstd, andreas schrempfd a czech academy of sciences, institute of theoretical and applied mechanics prosecká 809/76, 190 00 prague 9, czech republic b czech technical university in prague, faculty of transportation sciences, konviktská 20, 110 00 prague 1, czech republic c university of applied sciences upper austria in wels, computed tomography research group, stelzhamerstrasse 23, 4600 wels, austria d university of applied sciences upper austria in linz, research group for surgical simulators linz, garnisonstrasse 21, 4020 linz, austria ∗ corresponding author: sleichrt@itam.cas.cz abstract. the main goal of this study is to validate elementary mechanical parameters of a newly designed open-cell foam. the purpouse for investigating artificial material is to approach the properties of the human bone in the case of its adequate replacement. investigated material can be also used as an artificial bone to train surgical procedures and to improve the skills of the surgeons. four sets of the foam with different chemical composition were subjected to an uniaxial quasi-static loading to describe basic mechanical behaviour of these samples. based on these experiments, the stress-strain diagrams were created as a comparative tool including calculation of the effective young’s modulus. the acquired knowledges will be used as input parameters of a follow-up study aimed at describing the morphology of presented structures and their response to mechanical experiments. a distortion effect of porosity on the results is not considered in this study. keywords: polyurethane foam, artificial bone, uniaxial loading, digital image correlation. 1. introduction open cell foams are very effective lightweight structures with impressive load-bearing function. therefore there is wide range of application for this type of material in modern engineered products. moreover, this spongy-like structure is very close to natural foam materials hence it meets the presumption for the usage in bioengineering [1]. in this case different types of polyurethane foam synthesized as artificial cancellous bone was investigated from biomechanical point of view. the goal of the project is to find suitable material for artificial vertebrae filling with similar mechanical and structural parameters to real vertebra body [2]. resulting product will serve during the training of medical doctors in spine surgery simulator developed at fhoö linz [3]. based on results of previous studies [4], four different foam were synthetized employing various additives in base polyurethane matter. similarly to the natural cancellous bone, artificial structure exhibit complex porous microstructure composed of irregular polyhedral cells [5] which is influenced by the additives. bone aging and disease could be simulated the same way [6, 7]. set of pilot compression tests were performed on batch of samples prepared from all material to evaluate basic mechanical parameters and to test loading range for in-house developed loading device which is planned to be used for 4d computed tomography testing procedure as a next step in this project. 2. experimental procedure 2.1. specimen description two-component polyurethane resin (kaupo, spaichingen, germany) was used as the base material. depending on a mineral filler, four types of polyurethane foam were created (listed in tab. 1). each type of foam was provided by the research group for surgical simulators linz (university of applied sciences upper austria in linz, austria). a small amount of highly-concentrated colour pigments was added to distinguish individual sets of samples (depicted in fig. 1). the same aditives (in detail listed in [4]) were used in all cases, e.g. 1 % of tap water as a blowing agent. all the foam samples were delivered in cylindrical shape with thin other shell and cut by a band saw to smaller rollers with a diameter d = 19.70 ± 0.14 mm and height h = 15.42 ± 0.42 mm. 28 http://dx.doi.org/10.14311/app.2018.18.0028 http://ojs.cvut.cz/ojs/index.php/app vol. 18/2018 biomechanical characterization of polyurethane structures filler bulk density [g/dm3] 50% calcium carbonate (caco3) 1% h2o 3% stabilizer 416.69 ± 21.85 50% calcium phosphate (3ca3(po4)· cao) 1% h2o 3% stabilizer 362.15 ± 17.16 30% barium sulphate (baso4) 1% h2o 3% stabilizer 381.45 ± 36.07 without filler 1% h2o 3% stabilizer 467.33 ± 59.74 table 1. average bulk density of the samples figure 1. four types of samples – color-coded according to the filler 2.2. experimental setup experimental setup consisted of an in-house developed uni-axial loading device and optical setup. force was measured by high-durability u9b series loadcell with nominal loading capacity 1 kn (hbm, germany). the axial movement of a clamp was performed by an electrical stepper motor with a ball screw. displacementdriven tests were performed with 10 µm/s loading rate for 10 mm displacement (corresponding to ≈ 50 % deformation). force and displacement data were recorded at a sample rate of 5 hz. experiments were controlled by the modified linuxcnc software running on a real-time kernel [8]. because of the need to evaluate the deformation optically, the samples were captured by high-resolution ccd camera manta g-504bmanta (allied vision technologies gmbh) equipped with a telecentric lens. the optical setup was placed on a 3-axis stage equipped with micrometric screws for precise positioning. all of the images were captured by unique timestamp to synchronize the force log. during the measurement, place of view had to be enlightened by external led light source kl 2500 (shott, germany). complete equipment is depicted in fig. 2. 2.3. digital image correlation high precision of the deformation measurement is a necessary step for proper evaluation of the strainstress diagram. one of the common and widely used method is digital correlation image (dic). this optifigure 2. 1 camera, 2 positionable table, 3 led source, 4 loading device, 5 place for specimen figure 3. definition of the subset and offset cal method is based on a specific small area tracking subset, defined by a variable m. the region is then searched for in the wider neighborhood defined by a variable offset (shown in fig. 3). a necessary condition is to ensure a granular coating on the surface of each sample. primarily random pattern allows a tracking of these areas during the experiment and calculation of the displacement, necessary to calculate the deformation. updated lucas-kanade algorithm was used for this purpose [9, 10]. 3. results based on the obtained force-displacement results, the stress-strain curves were calculated using total dimensions of the samples (fig. 4 to 7). these results show dependency of the ultimate stress on the used filler in polyurethane foam. the stress-strain curves are very similar when using calcium phospate and barium sulphate filler. higher values of the monitored quantities are reached by the calcium carbonate filler. the highest values of the ultimate-stress and young’s modulus are achieved when no filler is used but these results are distorted by a thin shell of a high stiffness on the samples. young’s modulus was calculated using linear regression applied on the elastic part of the stress–strain 29 j. šleichrt, d. kytýř, k. pithartová et al. acta polytechnica ctu proceedings 0 0.5 1 1.5 2 2.5 0 0.1 0.2 0.3 0.4 0.5 s tr e ss [ m p a ] strain [-] samples a (50% calcium carbonate (caco3), 1% h2o and 3% stabilizer) a1, e = 51.357 mpa a3, e = 53.30 mpa a4, e = 46.49 mpa a5, e = 44.31 mpa figure 4. stress-strain curves of calcium carbonate filler 0 0.5 1 1.5 2 0 0.05 0.1 0.15 0.2 0.25 s tr e ss [ m p a ] strain [-] samples b (50% calcium phosphate (3ca3(po4) . cao), 1% h2o and 3% stabilizer) b1, e = 19.43 mpa b2, e = 20.72 mpa b3, e = 23.50 mpa b4, e = 27.53 mpa b5, e = 25.70 mpa figure 5. stress-strain curves of calcium phosphate filler 0 0.5 1 1.5 2 0 0.05 0.1 0.15 0.2 0.25 s tr e ss [ m p a ] strain [-] samples c (30% barium sulphate (bas04), 1% h2o and 3% stabilizer) c1, e = 37.891 mpa c2, e = 25.454 mpa c3, e = 24.953 mpa c4, e = 14.873 mpa c5, e = 21.867 mpa figure 6. stress-strain curves of barium sulphate filler diagrams. for the yield stress σy evaluation, christensen [11] second derivation criteria was used: σy = σ at ∣∣∣∣ d2σd�2 ∣∣∣∣ = max (1) for noise reduction significantly distorting resulting derived function rolling average filter computed from 0 1 2 3 4 5 0 0.05 0.1 0.15 0.2 0.25 s tr e ss [ m p a ] strain [-] samples d (no filler, 1% h2o and 3% stabilizer) d1, e = 50.27 mpa d2, e = 54.43 mpa d3, e = 104.81 mpa d4, e = 107.51 mpa d5, e = 128.16 mpa figure 7. stress-strain curves of no filler five values as: σ′na = σ′n−2 + σ′n−1 + σ′n + σ′n+1 + σ′n+2 5 (2) results of the effective young’s modulus of each of the sample’s batch are listed in tab. 2. filler effective young’smodulus [mpa] 50% calcium carbonate (caco3) 1% h2o 3% stabilizer 48.63 ± 4.37 50% calcium phosphate (3ca3(po4)· cao) 1% h2o 3% stabilizer 23.35 ± 3.89 30% barium sulphate (baso4) 1% h2o 3% stabilizer 25.01 ± 8.35 without filler 1% h2o 3% stabilizer 89.04 ± 34.72 table 2. average of the effective young’s modulus of the samples 4. conclusions all samples were subjected to the uniaxial compression test. digital image correlatioon method was used for evaluating the displacement during the measuring. due to very precise synchronization between captured images of the sample’s surface and force logs the stress-strain diagrams could be calculated and effective young’s modulus was calculated for each of the sample. high dispersion of the young’s modulus is described by the standard deviation. this indicates that the porosity or wall thickness vary in the volume of the sample. further investigation of the material, especially from morphological point of view, is necessary. for this reason, other samples 30 vol. 18/2018 biomechanical characterization of polyurethane structures will be subjected to an internal structure study using more advanced methods as 4d computed tomography testing procedure. acknowledgements the research has been supported by operational programme research, development and education in project competitiveness boost of the centre of excellence in vysočina region (cz.02.2.69/0.0/0.0/16_027/0008475), by european regional development fund in frame of the project com3d-xct (atcz38) in the interreg v-a austria czech republic programme and by institutional support rvo: 68378297. references [1] k. l. calvert, k. p. trumble, t. j. webster, l. a. kirkpatrick. characterization of commercial rigid polyurethane foams used as bone analogs for implant testing. journal of materials science: materials in medicine 21(5):1453–1461, 2010. doi:10.1007/s10856-010-4024-6. [2] t. fila, d. kytyr, p. zlamal, et al. high-resolution time-lapse tomography of rat vertebrae during compressive loading: deformation response analysis. journal of instrumentation 9(5), 2014. doi:10.1088/1748-0221/9/05/c05054. [3] b. esterer, j. razenbock, m. hollensteiner, et al. development of artificial tissue-like structures for a hybrid epidural anesthesia simulator. 2016 38th annual international conference of the ieee engineering in medicine and biology society (embc) pp. 2099–2102, 2016. doi:10.1109/embc.2016.7591142. [4] d. fuerst, s. senck, m. hollensteiner, et al. characterization of synthetic foam structures used to manufacture artificial vertebral trabecular bone. materials science and engineering c 2017. [5] m. s. thompson, i. d. mccarthy, l. lidgren, l. ryd. compressive and shear properties of commercially available polyurethane foams. j biomech eng 125(5):732–734, 2003. doi:10.1115/1.1614820. [6] h. chen, s. shoumura, s. emura, y. bunai. regional variations of vertebral trabecular bone microstructure with age and gender. osteoporosis international 19(10):1473–1483, 2008. doi:10.1007/s00198-008-0593-3. [7] h. chen, x. zhou, h. fujita, et al. age-related changes in trabecular and cortical bone microstructure. international journal of endocrinology (id 213234):9, 2013. doi:10.1155/2013/213234. [8] v. rada, t. fila, p. zlamal, et al. multi-channel control system for in-situ laboratory loading devices. acta polytechnica ctu proceedings 18:15–19 2018. doi:10.14311/app.2018.18.0015. [9] b. lucas, t. kanade. an iterative image registration technique with an application to stereo vision. proceedings of the 7th international joint conference on artificial intelligence 2:674–679, 1981. [10] i. jandejsek, j. valach, d. vavrik. optimization and calibration of digital image correlation method. proccedings ean 2010 pp. 121–126, 2010. [11] r. m. christensen. observations on the definition of yield stress. acta mechanica 196(3):239–244, 2008. doi:10.1007/s00707-007-0478-0. 31 http://dx.doi.org/10.1007/s10856-010-4024-6 http://dx.doi.org/10.1088/1748-0221/9/05/c05054 http://dx.doi.org/10.1109/embc.2016.7591142 http://dx.doi.org/10.1115/1.1614820 http://dx.doi.org/10.1007/s00198-008-0593-3 http://dx.doi.org/10.1155/2013/213234 http://dx.doi.org/10.14311/app.2018.18.0015 http://dx.doi.org/10.1007/s00707-007-0478-0 acta polytechnica ctu proceedings 18:28–31, 2018 1 introduction 2 experimental procedure 2.1 specimen description 2.2 experimental setup 2.3 digital image correlation 3 results 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.8.0024 acta polytechnica ctu proceedings 8:24–26, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app functionalization of 3d fibrous scaffolds prepared using centrifugal spinning with liposomes as a simple drug delivery system michala rampichováa, ∗, matej buzgoa, věra lukášováb, andrea míčkováa, b, karoliína vocetkováa, b, věra sovkováb, franco rustichellib, evžen amlera, b a uceeb, czech technical university in prague, bustehrad, czech republic b institute of experimental medicine of cas, prague, czech republic ∗ corresponding author: michala.rampichova@uceeb.cz abstract. 3d materials supporting cell adhesion, infiltration and proliferation are crucial for bone tissue engineering. to fulfill these expectations, it is beneficial to combine the 3d scaffold with drug delivery systems. release of the bioactive molecules directly in the site of the defect results in attraction of cells and their proliferation and differentiation in the required cell type. only properly differentiated cells produce proteins of extracellular matrix and can form new tissue. in the current study, we propose a simple drug delivery system composed of polycaprolactone fibers, prepared using centrifugal spinning, with adhered liposomes filled with platelet lysate. platelets contain wide range of growth factors and cytokines and can be used as their natural and autologous source. the scaffold was tested in vitro using mg-63 osteoblast-like cells. cell adhesion, proliferation and osteogenic differentiation were determined. three different concentrations of liposomes were used and compared to a scaffold with control liposomes filled only with pbs and to the plain polycaprolactone fibers. the results of the study showed that the platelet lysate stimulated cell proliferation. on the other hand, its effect on osteogenic differentiation was not proved. the combination of polymeric fibers with liposomes via simple adhesion was shown to be a promising scaffold for bone tissue engineering applications. in the future, it is necessary to focus on the stability of liposomes to prevent the burst release of the incorporated molecules. 1. introduction nowadays 3d materials supporting cell adhesion, infiltration and proliferation are needed in tissue engineering applications. these materials should serve not only as a scaffold, but also promote cell ingrowth and differentiation into the required cell type. therefore, it is beneficial to combine the scaffolding material with a drug delivery system. liposomes remain among the nanoparticles most widely used as carriers of active molecules [1]. nanofibers and microfibers are often used in tissue engineering, because they mimic the natural extracellular matrix (ecm) and enhance cell adhesion and proliferation [2]. centrifugal spinning uses centrifugal force to produce ultrafine fibers from melts and solutions [3], [4]. this method gives rise to fibers with submicron diameters and fluffy 3d structure. in the current study, pcl fibers prepared using the centrifugal spinning were combined with adhered liposomes. as a natural source of cytokines and growth factors the platelet lysate (pl) was used to boost the growth of mg-63 cells. 2. materials and methods 2.1. preparation of fibrous scaffold with adhered liposomes and its characterization fibrous meshes were prepared using centrifugal spinning device (cyclone 1000 l/m, fiberio, usa) from 40 % polycaprolactone (sigma aldrich) dissolved in chlorophorm:ethanol in volume ratio 9:1. orifice g30 and rotation speed 10000 rpm was used. unilamellar liposomes were prepared from l-α-phosphatidylcholine from egg yolk (sigmaaldrich) using the lipid film hydratation method followed by extrusion (1 µm pore size) [5]. liposomes were adhered to fibrous scaffolds for 2 hours and subsequently fibers were rinsed with pbs to wash out the non-adhered liposomes. five different samples were prepared: pcl fibers with maximal concentration of liposomes (25 mg/ml) (pcl/pl1); 4-fold diluted liposomes (pcl/pl2), and 10-fold diluted liposomes (pcl/pl3). plain pcl fibers (pcl) and pcl fibers with adhered liposomes contained pbs (pcl/pbs) were used as control samples. fibrous scaffolds with adhered liposomes were visualized using cryo-scanning electron microscopy (fei nova nanosem, cr) and size of liposomes was determined by mastersizer 3000. release of growth factors from liposomes with incor24 http://dx.doi.org/10.14311/app.2017.8.0024 http://ojs.cvut.cz/ojs/index.php/app vol. 8/2017 functionalization of 3d fibrous scaffolds figure 1. cryo-scanning electron microscopy of fibers prepared by centrifugal spinning techniques with adhered liposomes. detection of liposomes size distribution using dynamic light scattering (b) and detection of overall protein release (c). figure 2. mg-63 adhesion and proliferation on scaffolds measured using dna quantification (a) and cell metabolic activity measured using mts test (b). data were confirmed using confocal microscope. cell nuclei were stained using pi (red colour) and cell membranes using dioc6(3) (green). scale 200 µm, objective 10×. porated pl adhered on fibrous scaffolds was measured as a release of overall protein (quant-it protein assay kit, life technologies) and using elisa (tgf-β1, p-selectin; duoset®, r&d systems, usa). 2.2. in vitro test using mg-63 cells prepared samples were seeded with 3.1 × 104 mg63 cells/cm2. the metabolic activity of cells was measured using an mts assay (celltiter 96®aqueous one solution cell proliferation assay; promega). the proliferation of cells on the scaffold was measured from the amount of dna (quant-it™dsdna assay kit; life technologies). confocal microscopy (zeiss lsm duo) was used to visualize cell adhesion on the scaffolds. samples were stained using fluorescent probe 3,3’diethyloxacarbocyanine iodide (dioc6(3), invitrogen, 1 µg/ml in pbs) and propidium iodide (pi; 5 µg/ml in pbs). osteogenic differentiation was detected as synthesis and production of proteins typical for bone ecm, type i collagen and osteocalcin using real-time pcr (light cycler 480 ii; roche, switzerland) and immunofluorescence staining (osteocalcin ocg3, abcam, uk, af488 secondary antibody). results were evaluated statistically using one way anova. 3. results and discussion 3.1. liposome adhesion on fibers and release of protein content figure 1 shows adhesion of liposomes on the fibers. liposome size distribution measurement showed that 67 % of liposomes were in range 0.46 0.77 µm (fig 1b). the pattern of release was similar for tested proteins. there was visible burst release from liposomes after the first 15 minutes (fig 1c). the strongest release was detected in pcl/pl1 sample with the most liposomes. small release was visible on pcl/pl1 and pcl/pl2 samples 24 hours after liposome adhesion. the burst release of proteins was caused by low stability of the liposomes. 25 m. rampichová, m. buzgo, v. lukášová et al. acta polytechnica ctu proceedings 3.2. cell viability, proliferation and osteogenic differentiation on scaffolds cells adhered similarly on all scaffolds. cell proliferation was mostly visible on pcl/pl3 and pcl/pl2 respectively. after 14 days of cultivation, dna amount was higher on all samples containing thrombocyte lysate (pcl/pl1 3) compared to the sample with control liposomes with pbs (pcl/pbs) (fig 2a). cell viability was comparable for all samples (fig 2b). data were confirmed by confocal microscopy. cells adhered well on all samples and were uniformly distributed. cells proliferated till day 7, when confluent layers of cells were visible on pcl/pl1 3 (fig 2e-g). decrease in cell number was seen on day 14. detection of osteogenic markers was very low and did not show any differences between experimental and control samples (data not shown). positive influence of pl was shown in following days. interestingly, positive effect was more visible in samples with 4and 10-fold diluted liposomes (pcl/pl2 and pcl/pl3) and thus pl platelets contain a wide spectrum of growth factors and cytokines with different effect. in our study positive effect on cell proliferation was detected. on the other hand, no effect was shown on osteogenic differentiation. positive effect of pl was limited by its burst release from the liposomes. surface adhesion is simple and variable method of functionalization of the surface of fibers. one of the possibilities how to prolong the release is stabilization of liposomes using peg. another strategy is to incorporate bioactive molecules inside the fibers. up to date, only limited number of studies report using of forcespinning fibers as a drug delivery system [6]. 4. conclusion combination of 3d fibers prepared using centrifugal spinning and liposome adhesion can serve as simple and variable drug delivery system. in the future study we need to focus on liposome stability to prolong the release of incorporated molecules. acknowledgements this research was supported by the czech science foundation grant no. 15-15697s. authors thank to dr. benada from institute of microbiology of the cas for cryo-sem images. references [1] a. lowery, h. onishko, d. e. hallahan, z. han. tumor-targeted delivery of liposome-encapsulated doxorubicin by use of a peptide that selectively binds to irradiated tumors. journal of controlled release 150(1):117 – 124, 2011. doi:http://dx.doi.org/10.1016/j.jconrel.2010.11.006. [2] m. rampichová, j. chvojka, m. buzgo, et al. elastic three-dimensional poly (�-caprolactone) nanofibre scaffold enhances migration, proliferation and osteogenic differentiation of mesenchymal stem cells. cell proliferation 46(1):23–37, 2013. doi:10.1111/cpr.12001. [3] a. m. loordhuswamy, v. r. krishnaswamy, p. s. korrapati, et al. fabrication of highly aligned fibrous scaffolds for tissue regeneration by centrifugal spinning technology. materials science and engineering: c 42:799 – 807, 2014. doi:http://dx.doi.org/10.1016/j.msec.2014.06.011. [4] m. rampichová, m. buzgo, j. chvojka, et al. cell penetration to nanofibrous scaffolds: forcespinning®, an alternative approach for fabricating 3d nanofibers. cell adhesion & migration 8(1):36 – 41, 2014. doi:10.4161/cam.27477. [5] l. d. mayer, m. b. bally, m. j. hope, p. r. cullis. techniques for encapsulating bioactive agents into liposomes. chemistry and physics of lipids 40(2):333 – 345, 1986. doi:http://dx.doi.org/10.1016/0009-3084(86)90077-0. [6] l. a. mary, t. senthilram, s. suganya, et al. centrifugal spun ultrafine fibrous web as a potential drug delivery vehicle. express polymer letters 7(3):238 – 248, 2013. doi:10.3144/expresspolymlett.2013.22. 26 http://dx.doi.org/http://dx.doi.org/10.1016/j.jconrel.2010.11.006 http://dx.doi.org/10.1111/cpr.12001 http://dx.doi.org/http://dx.doi.org/10.1016/j.msec.2014.06.011 http://dx.doi.org/10.4161/cam.27477 http://dx.doi.org/http://dx.doi.org/10.1016/0009-3084(86)90077-0 http://dx.doi.org/10.3144/expresspolymlett.2013.22 acta polytechnica ctu proceedings 8:24–26, 2017 1 introduction 2 materials and methods 2.1 preparation of fibrous scaffold with adhered liposomes and its characterization 2.2 in vitro test using mg-63 cells 3 results and discussion 3.1 liposome adhesion on fibers and release of protein content 3.2 cell viability, proliferation and osteogenic differentiation on scaffolds 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0080 acta polytechnica ctu proceedings 11:80–85, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app cant excess and cant deficiency evaluation in relation to train speed jan valehrach∗, tomáš říha, erik dušek brno university of technology, faculty of civil engineering, institute of railway structures and constructions, veveří 331/95, 60200 brno, czech republic ∗ corresponding author: valehrach.j@fce.vutbr.cz abstract. the paper deals with the evaluation of measured speeds of the vehicles on the chosen sections of the railway network in the czech republic. in connection with the phased implementation of the avv system (automatic train operation) and the european etcs system, the dispersion of the train speeds will be reduced. measured data were used to assess the weighted average of the cant excess or cant deficiency in relation to the rail defects. within the framework of the research, the speeds were statistically evaluated covering more aspects like types and sorts of the trains, their operators, daytime and more. keywords: speed/velocity, track, train, cant excess, cant deficiency. 1. introduction in the present time speeds of trains passing through the certain place vary. the variability of the speeds depends on more factors, such as train composition and category, actual traffic situation, daytime and the location. in connection with the implementation of new systems for train controlling and commanding the speed profile of the trains will likely become more efficient. by author’s opinion, this will eventually lead to the consolidation of speeds into a narrow range, or eventually into more ranges based on above mentioned factors. the knowledge of the speeds is necessary for a number of fields in railway industry. some of the applications, where the knowledge of speed is critical, are the studies [1–3]. this article focuses on the influence of speed on development of defects on superstructure by studying speed ranges in certain places of the railway infrastructure. as a starting point for the study an expected relation between the appearance of the long-pitch corrugation in curves of small radii and the speed range of the passing vehicles was considered. authors believe that the speed range in certain places of the infrastructure can affect the functionality of the superstructure and under specific circumstances the inappropriate speed range could result in the development of the defects of the rails. by minimizing inconvenient combinations of track parameters and the speed ranges the positive effect will likely appear, which in some extents gives us the potentiality for the optimization. one of the possible parameters, influencing the appearance of the long pitch corrugation in curves of small radii is considered to be cant excess or cant deficiency. in standard conditions, where negative effects on infrastructure do not take place, both values are designed with the design of the track itself. in such case, the designer uses some initial expectations of future traffic composition to calculate cant values. then uses (still only as an expectation) the track speed and calculates the values of cant excess or cant deficiency. nevertheless, this method could not take in account the real speeds of vehicles, which may be far different from the expectations. this lack of feedback could result in some negative effects on infrastructure, and one of them is believed to be the long pitch corrugation defect. this work focuses on evaluation of the cant excess or cant deficiency values obtained through the speed measurements while the relation to long pitch corrugation is studied. 2. methods of measurement and evaluation 2.1. location selection the appearance of the long pitch corrugation was the basis for the selection of the locations for measurements. extensive measurements took place in first track of the line from brno to česká třebová between stations adamov and brno-maloměřice. four locations were selected as follows: • line no. 326 (brno – česká třebová) near babice nad svitavou, km 169.0, in approximately 200 metres the first turnout of the marshalling yard brnomaloměřice follows; • line no. 326 (brno – česká třebová) near babice nad svitavou, km 167.0, the position of the automatic block signal; • line no. 326 (brno – česká třebová) near former branching station hady, km 161.7; • line no. 326 (brno – česká třebová) near former branching station hady, km 162.8, currently the babice stop. 80 http://dx.doi.org/10.14311/app.2017.11.0080 http://ojs.cvut.cz/ojs/index.php/app vol. 11/2017 cant excess and cant deficiency evaluation figure 1. the locations positioning of speed recording device (red points) and automatic image recorder (blue points). in all four locations the significant long pitch corrugation defect occurs. 2.2. measurement method all the measurements were continual and lasted several days. measurements were conducted in pairs, since there are two recording devices at disposal. the pair includes the measurements in km 169.0 and in km 167.0, the second pair consists of measurements in km 161.7 and in km 162.8. by pairing it is possible to get the speed of one particular train in two locations. following data were recorded for each passing train: • speed; • image record; • date and time of the recorded event. as a measuring instrument for recording speeds and directions of the trains were used drs devices (frequency modulation radar units). the main advantage of the device is the possibility of the quick on site installation and measuring without any external energy supply due to device batteries. on the other hand, the device tends to record more than one speed of the train, partially due to the response time and partially due to the length of the trains, which exceeds the length of the vehicles for which was the device initially built1. thus, longer trains inflict more measurements and more records. such sets of records then have to be sorted out by hand, which is very challenging work. the aim is to choose only the speed record which corresponds with the speed of a head of a train. mostly, the first record is chosen, but in some cases careful consideration has to take place. 2.3. evaluation following information about the particular train is then searched out based on the image record, train traffic diagram, timetable, train compositions list and others: • category (freight train, express, passenger train and more – see below); • train number; • composition of the train or train set, types of couches; • train mass; • delay. despite of a great amount of endeavour some trains could not be identified satisfactory. also some measurements had to be edited out, mainly the ones anticipated faulty. at the end, in every of all four locations at least 200 trains were evaluated. 2.4. calculation method a weighted average of all speed measurements in a location was calculated, when the mass of each train was taken into consideration [4]. the masses of the trains for the purposes of the weighted averages have been obtained by two ways. the weight of passenger trains was taken from the passenger carriages charts and from the train compositions list. the mass of the freight trains could be found out from the mass normative (which is the maximum allowable mass of the whole train). the problem is that the actual load on the train is not known and thus it is not possible to know the right mass of the train. therefore, the masses of the freight trains were estimated with the value 1000 tonnes2. this approach brings the significant amount of uncertainty to the result, but on the other hand, according to higher order masses of the freight trains in comparison to the passenger trains, the sensitivity of the result to the changes of the masses is quite small. then for the calculation of the cant excess and the cant deficiency the following scheme was used [5]. values of radii and cant for every location were taken from superstructure plotted outline [6]. 1used devices were initially designed for recording of road vehicles not exceeding the length of 18 metres. 2despite of a fact that this simplification seems to be very disproportional, the results later imply, that the influence of a train masses on the cant excess and cant deficiency is not as strong as was expected. in the present time authors believe, that the influence of the train mass could be negligible. by all means, this still certainly has to be researched more thoroughly. 81 j. valehrach, t. říha, e. dušek acta polytechnica ctu proceedings assuming a formula for ideal speed of trains v = √ d × r 11, 8 (1) the cant excess of a single train then applies ei = d − 11, 8 × v 2i r (2) and analogically for the cant deficiency ii = 11, 8 × v 2i r − d (3) ideally, the effects of cant excess and deficiency should be balanced, especially with respect to train masses. however, in general in some locations cant excess could markedly dominate. that happens for example in the vicinity of stations, where the trains are accelerating and have not yet reached the track speed or where the greatest amount of the track load consists of slow moving trains (usually freight trains). in such locations this leads to the overload of the railway superstructure and consequently results in increased occurrence probability of the defects. 3. outputs 3.1. speed ranges the first look at the cleaned data inserted into the charts is promising. the charts below show all four evaluated locations. the speed ranges are apparent, and also the variability could be easily distinguished. the width of the ranges could be identified as well. the width of the speed range has to be considered separately from the variability. the variability describes more the whole dispersion of the measured values, whileas the speed range could be in most cases identified not only once (namely in first chart there could be identified the upper speed range between 65 and 80 and the down speed range between 45 and 60). the speed range however should include the category of trains, as described below in figure 3 and thus it is more complicated to identify it soundly. some of the records are separated from the speed ranges. these records are considered as anomalies; since they are isolated, the passes of the trains corresponding to these records cannot have any significant influence on the infrastructure3. for these purposes, these records are in further analysis neglected. the following table differentiates the speeds by category of the train, showing the average values. 3.2. cant excess and deficiency values for every single train the cant excess or cant deficiency was calculated. in the following charts also the categories of the trains were distinguished, showing clearly the anomalies in the traffic. 3in general, exemptions to this statement could be found, such as special transports of overweighed cargo or other similar uncommon traffic. as it could be seen in the figure 3, the highest cant deficiency values (as high as 130 mm) are used by passing trains in location 162.8. however, this happens to be the location with the least cant of all monitored locations. on the contrary, the least cant deficiency values are used by passing trains in location 169.0. in this case the track alignment consists of compound curve of a radius 375 m with cant 140 mm in first part and radius 382 and cant 100 mm in second part of the curve. by the end of the curve could be found the train stop. measuring section was located in first part of this curve. as a result of evaluation according to chapter 2.4 just one value of the weighted cant deficiency was calculated. in case of our locations, cant excess did not take place. this value could be used for specifying the influence of all the trains passing through certain location on the infrastructure [4]. in case of a need of a closer in-depth analysis, it is necessary to divided the trains into more groups or categories and to calculate the weighted averages afterwards. 3.3. selected categories the categories were selected based on the categories of the passenger trains (also private companies were taken into account). in the future it could be expected, that the speeds will consolidate within these categories rather than by the construction and design properties of the trains. for every category the weighted cant excess (formula 4) or cant deficiency (formula 5) were calculated by e = ∑n i mi × ei∑n i mi (4) i = ∑n i mi × ii∑n i mi (5) results calculated using formulas 4 and 5 are displayed in table 2. when comparing the weighted cant excess or cant deficiency for train categories with the weighted value for all the trains according to table 3 it could be stated, that at least for the rough estimate for the purposes of setting curves against each other is this value conclusive enough. 3.4. train speed/velocity and long-pitch corrugation in a framework of the research also the rate of long pitch corrugation development was monitored in locations 161.7 and 162.8. the corrugation tends to develop quicker in the curve with 162.8, which is the curve where higher values of cant deficiency were found [7]. 82 vol. 11/2017 cant excess and cant deficiency evaluation figure 2. first look at cleaned data from all locations. based on this finding and also using experience from the preceding investigation it is to expect that in case of curves with locations 167.0 and 169.0 the long pitch corrugation will be developing faster in curve with 167.0. 4. conclusions the research was aimed to confirm or disprove the possibility, that cant excess or cant deficiency could influence the development of certain defects of the railway superstructure. in order to be able to calculate the actual values of cant excess or cant deficiency it was necessary to measure speed of passing vehicles. four locations were chosen, both for comparison and for variability monitoring. the speeds were measured with the recording devices and the trains were divided into categories of the similar speed range. after that, the calculation of the cant deficiency or cant excess was executed with double results. first set of results uses as a basis all the speed measurements across all the train categories. second set of results respects the averages in categories detached. the results show, that even the simpler first method could bring satisfactory results. further research should focus on the long-term monitoring of train speeds in order to be able to calculate cant deficiency or cant excess more precise. also more emphasis should be put into the differentiation of masses of the trains. acknowledgements research reported in this paper was supported by competence centres program of technology agency of the czech republic (ta cr), project centre for effective and sustainable transport infrastructure (no. te01020168), student’s grant tender of the faculty of the civil engineering of the brno university of technology (fast-j-15-2860) and railway infrastructure administration, state organization (sždc s. o.). references [1] c. esveld. c. modern railway track. 2nd ed. zaltbommel: mrt-productions, 2001. isbn 90-800-3243-3. [2] j. smutny, i. vukusic, v. tomandl, et al. analýza dynamických účinků v srdcovkové části výhybek, stavební obzor. stavební obzor, thákurova 7, 166 29 praha 6, 2013. issn 1210-4027. [3] p. guziur. design parameters of buffer stops. acta polytechnica ctu proceedings. 2016. issn: 2336-5382. 83 j. valehrach, t. říha, e. dušek acta polytechnica ctu proceedings figure 3. the comparison of the cant excess and cant deficiency in locations (cant deficiency [+] / cant excess [-]). 84 vol. 11/2017 cant excess and cant deficiency evaluation km average speed v [kph] for the train categoryeurocity train1 express train2 euronight train3 fast train4 fast train suspended5 passenger train6 railjet train7 regiojet train8 freight train other trains9 161.7 72.93 67.80 73.00 68.87 72.13 68.47 71.13 72.80 58.27 53.00 162.8 73.50 73.40 73.00 69.57 72.75 70.09 71.60 73.80 61.64 60.67 167.0 75.18 75.18 75.18 75.18 75.18 75.18 75.18 75.18 75.18 75.18 169.0 78.68 78.68 78.68 78.68 78.68 78.68 78.68 78.68 78.68 78.68 1 the highest category of passenger trains, calling only at major stations not less than 60 km away from each other, usually consisting of the locomotive and six to eleven coaches (hereinafter ec). 2 the category of passenger trains, calling only at major stations not less than 60 km away from each other, usually consisting of the locomotive and three to six coaches (hereinafter ex). 3 the highest category of passenger trains running during nigh time, calling only at major stations not less than 60 km away from each other, usually consisting of the locomotive and more than ten coaches (hereinafter en). 4 the category of passenger trains below express trains, calling at major stations and junctions, usually represented by the train set interpanter (hereinafter r). 5 the category of passenger trains not calling at all stops, usually consisting of the locomotive and three to six coaches (hereinafter sp). 6 the usual passenger train calling at all stops. the train composition differs, both locomotive with coaches and train sets are used to cover this category (hereinafter os). 7 the same category of passenger trains as the eurocity trains, consisting exclusively of the locomotive and train set siemens viaggio comfort (hereinafter railjet). 8 the same category of passenger trains as the eurocity trains, consisting exclusively of the locomotive and six to eleven coaches (hereinafter regiojet). 9 covers all other categories of trains, or trains which do not have specific category, for example maintenance machines etc. table 1. average speeds of the displayed train categories. km weighted cant deficiency [+] / cant excess [-] in millimetres for the categories of trainseurocity train express train euronight train fast train fast train suspended passenger train railjet train regiojet train freight train other trains 161.7 100 82 100 72 92 74 89 99 21 -4 162.8 122 124 118 93 114 98 110 124 52 42 167.0 90 98 96 90 83 78 94 95 58 86 169.0 56 57 49 56 57 34 54 55 8 51 table 2. cant deficiency and cant excess values. km average speed v [kph] for all trains 161.7 67.84 162.8 70.00 167.0 75.18 169.0 78.68 table 3. average speeds of all trains. [4] j. valehrach, j. smid. monitoring of track sections with long-pitch corrugation, proc. int. conf. on interoperability of railway transport – iricon 2016. czech technical university in prague, 2016. isbn 978-80-01-05939-5. [5] c. . 6360-1. geometrical characteristics of railway tracks – part 1. layout. 2008. [6] s. o. szdc. superstructure plotted outline, internal km all trains 161.7 57.1 162.8 83.1 167.0 73.7 169.0 27.7 table 4. cant deficiency values (not rounded). information system output. valid for 2017-04. brno, 2017. [7] j. valehrach. long-pitch corrugation – monitored track sections, proc. int. conf. on young scientist. technical university of kosice, faculty of civil engineering, kosice, 2016. isbn 978-80-553-2537-8. 85 acta polytechnica ctu proceedings 11:80–85, 2017 1 introduction 2 methods of measurement and evaluation 2.1 location selection 2.2 measurement method 2.3 evaluation 2.4 calculation method 3 outputs 3.1 speed ranges 3.2 cant excess and deficiency values 3.3 selected categories 3.4 train speed/velocity and long-pitch corrugation 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0080 acta polytechnica ctu proceedings 13:80–84, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app adaptive quasicontinuum simulation of elastic-brittle disordered lattices karel mikeš∗, milan jirásek czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: karel.mikes.1@fsv.cvut.cz abstract. the quasicontinuum (qc) method is a computational technique that can efficiently handle atomistic lattices by combining continuum and atomistic approaches. in this work, the qc method is combined with an adaptive algorithm, to obtain correct predictions of crack trajectories in failure simulations. numerical simulations of crack propagation in elastic-brittle disordered lattices are performed for a two-dimensional example. the obtained results are compared with the fully resolved particle model. it is shown that the adaptive qc simulation provides a significant reduction of the computational demand. at the same time, the macroscopic crack trajectories and the shape of the force-displacement diagram are very well captured. keywords: quasicontinuum method, adaptive crack propagation, elastic-brittle lattice model. 1. introduction discrete particle models can effectively capture complex material responses, especially localized phenomena such as damage or plastic softening. the main disadvantage of particle-based approaches is that a huge number of particles is needed to describe the response of large-scale physically relevant models. this results in an extensive system of equations, which is expensive to solve. furthermore, the process of assembling this system is also computationally expensive because all discrete connections must be individually taken into account. a quasicontinuum (qc) based method can remove both of these disadvantages by combining continuum and atomistic approaches. in order to simulate crack propagation, the qc method needs to be combined with an adaptive algorithm that allows crack growth in arbitrary directions and initialization of new cracks. in this work, the material is represented by particles interacting via elastic-brittle links forming a disordered two-dimensional lattice. only axial interaction between particles is considered and the behavior of links is assumed to be perfectly elastic-brittle, with link breakage occurring at a critical level of tensile strain. 2. qc method the qc method can efficiently handle high-resolution particle models by combining continuum and discrete approaches. this method was originally proposed in 1996 [1] for regular atomistic lattices with long-range interactions. since that time, the qc method has been widely used and extended to applications for a variety of materials represented by regular lattices [2]. an extension of the qc method to irregular lattices has recently been developed by the authors [3]. the main idea of the qc method is to reduce the number of degrees of freedom (dofs) and the associated computational cost without losing the exact atomistic description in regions of interest. therefore, two types of regions in the investigated domain are considered. in regions of high interest, the pure particle approach is used and all particles carry their own independent dofs. by contrast, in regions of low interest, continuum assumptions are applied and the computational model is significantly simplified. the procedure that results from the qc method can be briefly presented in the following three steps: (1.) interpolation (2.) summation (3.) adaptivity 2.1. interpolation interpolation of dofs is used in regions of low interest. to simplify the full particle model, only a small subset of particles is selected to represent the entire system. these so-called repnodes serve as the nodes of a background triangular mesh that is used to interpolate the dofs of other particles in regions of low interest. on the other hand, in regions of high interest, all particles are selected as repnodes to provide the exact particle representation. this interpolation leads to a significant reduction of the number of dofs without losing the exact particle description in regions where a high resolution is needed. 2.2. summation the interpolation provides a significant reduction of the number of dofs but all particles still need to be visited to construct the system of governing equations, which makes the process computationally expensive. 80 http://dx.doi.org/10.14311/app.2017.13.0080 http://ojs.cvut.cz/ojs/index.php/app vol. 13/2017 adaptive quasicontinuum simulation of disordered lattices the summation rule can be applied in order to eliminate the requirement of visiting all particles during the assembly of the system. if the summation rule is adopted, the contribution of all particles in each interpolation triangle is estimated based on sampling of the links that surround one single particle and properly scaling the contribution. this makes the computational process faster but some problems can occur on the interface between regions of high and low interest. because of the interpolation and the summation, the deformation is considered as constant within each interpolation element in the regions of low interest while the deformations of individual links in the regions of high interest are evaluated exactly. consequently, forces of nonphysical character, called the ghost forces, appear on the interface between the regions of low and high interest [4]. in this work, the summation procedure is realized by a homogenization of links contributing to the interpolation elements. to eliminate the ghost forces, some specific links are selected to be processed exactly, in order to capture the interface between the fully resolved and interpolated domains. 2.3. adaptivity adaptivity allows to adjust the regions of low and high interest during the simulation process. the type of region can be changed by adding repnodes before each step. an appropriate modification of the regions of high interest leads to a substantial increase of accuracy. moreover, in specific cases such as simulation of crack propagation or damage evolution, adaptivity is necessary in order to represent the correct physical behavior. after the insertion of new repnodes, a new triangulation of the interpolation mesh can be constructed, in order to improve the accuracy of interpolation and summation in areas of low interest. 3. qc-based approaches in this section, the exact approach and two approaches based on the idea of qc with different levels of simplification are introduced. 3.1. fully resolved approach this approach does not use any simplification. every single particle represents a node with independent dofs. all links are taken into account explicitly and contribute directly to the stiffness matrix. consequently, this approach provides the “exact” result, which is used as a reference solution for evaluation of accuracy and efficiency of the following, simplified approaches. 3.2. qc approach with interpolation in this approach, only the interpolation rule is used to simplify the full particle model. in the regions of high interest, all particles are selected as repnodes. by contrast, in the area of low interest, only the particles forming the vertices of interpolation elements are selected as repnodes. all remaining particles are kept in the model and their dofs are linearly interpolated by using the underlying interpolation mesh. such particles are called hanging nodes because their dofs are not independent but are “hanging” on appropriate repnodes. all link connections are considered explicitly in the whole domain. 3.3. qc approach with interpolation and homogenization in this approach, 2d finite elements are used not only to interpolate dofs but also to replace the stiffness that corresponds to the microstructure. consequently, a substantial number of links and hanging nodes can be removed from the particle model. material properties of 2d elements are identified by homogenization of the effective elastic stiffness tensor de representing the microstructure of the links. applying the equivalence of the overall virtual work expressed for the continuous material and for the microstructure represented by discrete links, the following formula for the effective elastic stiffness tensor can be derived: de = 1 v nt∑ i=1 lieaini ⊗ ni ⊗ ni ⊗ ni (1) here, ni is the unit vector specifying the direction of the given link, e is the young modulus, a is the cross-sectional area and l is the length of the link. the summation is taken over nt links occupying volume v . the effective stiffness tensor de can be evaluated globally for all elements, or locally for each element separately. afterwards, the material parameters of 2d elements can be identified as isotropic, orthotropic, or arbitrarily anisotropic. different homogenization procedures for disordered lattices with normal interactions are described in [3]. in this work, only the most accurate homogenization rule with a local anisotropic stiffness tensor and an arbitrarily anisotropic material is used. evaluation of the effective material stiffness tensor given by (1) is done for each element separately. the stiffness tensor of each element is obtained only from the contributions of the parts of the links that are located in that particular element. 4. adaptive algorithm adaptivity is one of the key ingredients of the qc method. the area of high interest and the geometry of interpolation mesh can be arbitrarily changed during the simulation to optimize the accuracy and computational costs. furthermore, for qc simulation of materials with softening or damage, an adaptive 81 karel mikeš, milan jirásek acta polytechnica ctu proceedings algorithm is required to be able to predict the exact failure mechanism. in this paper, it is assumed that cracked links tend to form a macroscopic crack. the qc adaptive algorithm is designed to be able to simulate the growth of existing macroscopic crack as well as the initialization of new cracks. in each simulation step, this algorithm is able to change both the area of high interest and the interpolation elements in the area of low interest. the overall algorithm of an adaptive cq simulation is organized as follows: algorithm 1 adaptive qc simulation 1: load input geometry 2: generate interpolation mesh 3: calculate local stiffness tensors and material parameters 4: assemble stiffness matrix 5: assemble load vector 6: repeat 7: solve one elastic step 8: break link with maximal strain 9: update area of high interest 10: update interpolation mesh 11: update local stiffness tensors and material parameters 12: update stiffness matrix and load vector 13: until <stopping criterion> 14: postprocessing 4.1. area of high interest update the change of the area of high interest is realized around newly cracked links to allow crack growth. when a new link is broken, all particles within a given updating radius are labeled as repnodes. this update is realized only if the breaking link is located not too far from existing macroscopic cracks. therefore, a number of previously cracked links within a checking radius is evaluated and the update is realized only if a given critical number of previously cracked links is reached. consequently, the area of high interest is updated frequently but not necessarily in all steps. this allows to account not only for crack branching but also for nucleation of new macroscopic cracks in different locations. this procedure frequently requires searching for particles within a given radius, which can be timeconsuming for large sets of particles. therefore, it is convenient to store particles in an octree structure that enables fast spatial searching. 4.2. interpolation mesh update if the area of high interest has changed, it seems natural to change the geometry of the interpolation mesh as well to provide a better shape of interpolation elements at the interface between areas of low and high interest. however, this change is extremely expensive, as explained in section 4.3, and does not bring any significant improvement of accuracy. therefore, it turns out to be more effective to use relatively small elements with fixed geometry instead of changing the size of elements during the simulation. only the elements that are completely covered by the extended area of high interest become useless and can be removed from the model. if the interpolation elements are not changed, the updating radius must be larger than the size of interpolation elements to ensure that the nodes of one interpolation element are not located on the opposite sides of the crack (otherwise, special finite element techniques such as xfem must be used). 4.3. stiffness tensors update generally, there are three reasons why stiffness tensors must be updated. (1.) breakage of a link (2.) change of a region of high interest (3.) change of the interpolation mesh the breakage of one link occurs in each simulation step. the resulting update of stiffness tensors can be simply realized by subtracting the stiffness contributions of broken links from all corresponding stiffness tensors. the change of the area of high interest is realized by adding new repnodes. consequently, the type of ending particles of some links is changed. all links changed to type repnode–repnode and some specific links changed to type repnode–hangingnode need to be solved explicitly from this moment. all such links must be identified and their stiffness influence is subtracted in the same way as for broken links. the change of the geometry of interpolation mesh requires a recalculation of local stiffness tensors. this recalculation is extremely demanding because all links must be visited. even if the mesh is changed only locally near a newly created area of high interest and only some elements are affected, it is not easy to identify which links contribute to these elements and this step significantly slows down the simulation time. the recalculation of stiffness can be sped up by storing the information of element contributions during the assemblage of stiffness tensors at the beginning of the simulation. but this can be done only at the price of additional memory demands, which may become critical for large numbers of links. 5. results to assess the efficiency and accuracy, the presented adaptive qc algorithm has been used to simulate failure in two dimensions. the l-shaped specimen with dimensions 100 × 100 mm (which may correspond to a joint of a frame) is fixed at the bottom section and loaded by prescribed vertical displacements imposed at the right end section; see figure 1. as a result of this loading, the non-convex corner is opened 82 vol. 13/2017 adaptive quasicontinuum simulation of disordered lattices 0 20 40 60 80 100 0 10 20 30 40 50 60 70 80 90 100 ↑ ↑ ↑ ↑ ↑ ↑ 0 20 40 60 80 100 0 10 20 30 40 50 60 70 80 90 100 ↑ ↑ ↑ ↑ ↑ ↑ 0 20 40 60 80 100 0 10 20 30 40 50 60 70 80 90 100 ↑ ↑ ↑ ↑ ↑ ↑ figure 1. interpolation elements and area of high interest (black) and broken links (red) in the first (left), 100-th (middle) and the last step of simulation (right). 0 20 40 60 80 100 0 10 20 30 40 50 60 70 80 90 100 figure 2. broken links computed with exact particle approach a0 (black) and with qc approach with local anisotropic homogenization a3 (red). and cracking of links is expected to start in this region. therefore, a small initial area of high interest is prescribed around this corner; see figure 1 (left). the random microstructure is generated with a density that corresponds to 67 particles along the short edge, leading to a total of 19,000 particles connected by 64,064 links. basic characteristics of models used by full particle and qc simplified simulations are listed in table 1. material parameters are considered to be the same for all links. as expected, cracking starts in the non-convex corner. duringcrack growth, the area of high interest is increased around newly broken links. interpolation elements completely covered by the area of high interest are removed from the model, but the geometry of the remaining interpolation elements is not changed. the areas of high interest in different loading steps are depicted in figure 1. overall 240 simulation steps 0 1 2 3 4 0 5 10 15 20 25 30 35 40 45 displacement [mm] f o rc e [ n ] exact particle approach a0 qc simplified approach a3 figure 3. force-displacement diagrams computed with exact particle approach a0 (black) and with qc approach with local anisotropic homogenization a3 (red). a0 a1 a3 particles 19,000 19,000 2,199 links 64,064 64,064 2,929 elements 0 1,542 1,542 repnodes 19,000 1,665 1,665 hnaging nodes 0 17,335 534 dofs 37,799 3,129 3,129 table 1. numbers of particles, links, elements, repnodes, hanging nodes and unknown dofs for exact approach (a0), qc approaches with interpolation only (a1) and qc approach with local anisotropic homogenization (a3). are performed until the specimen is almost completely split; see figure 1 (right). three different levels of simplification are used and compared. • full (exact) particle approach (a0). • qc simplified approach with interpolation only (a1). 83 karel mikeš, milan jirásek acta polytechnica ctu proceedings • qc simplified approach with interpolation and local anisotropic homogenization (a3). the results obtained for both qc simplified approaches are almost equivalent. both approaches predict the same broken links in the same order. generally, approaches involving homogenization tend to a stiffer response in comparison with the interpolation approach. in this case, the influence of homogenization is negligible and the maximum difference in relative stiffness is just 0.24 %. in comparison with the exact approach, the qc approaches give a slightly stiffer response. the relative stiffness error of the initial elastic branch is 11.8 %. the maximum relative stiffness error is below 18 % during almost the whole simulation. only in the last six steps (when the specimen is almost completely split) the error increases up to 25 %. the prediction of cracked links is not perfectly exact, but more than 90 % of links are broken correctly and the difference in the final shape of the macroscopic crack is almost negligible; see figure 2. the simplified results are also very accurate in terms of the force-displacement diagram. despite a stiffer elastic response, the trend of the unloading branch is very well captured and the shape of the qc simplified force-displacement diagram resembles the exact solution; see figure 3. times consumed by individual parts of the simulation for various approaches are listed in table 2. the initialization procedures are realized just once at the beginning of the simulation. on the other hand, the repeated procedures are realized in each simulation step. for qc approach a1, with interpolation only, the realization of qc simplification is relatively fast, but the initial assemblage of the stiffness matrix and load vector is extremely slow because of a huge number of hanging nodes with interpolated dofs. by contrast, qc approach a3, involving local homogenization, needs a relatively long time to realize qc simplification including local effective stiffness tensors calculation. however, the initial assemblage is very fast because the number of links and hanging nodes is significantly reduced. for both qc approaches, a long extra time is needed only at the beginning of the simulation. this initial time becomes negligible in simulations with a large number of repeated steps. the most important fact is that each simulation step of the simplified approach is twice as fast for a1 and almost 19 times faster for a3 in comparison with the full particle model. 6. conclusions this paper described a qc adaptive algorithm with automatic changes of area of high interest designed for simulation of crack propagation in disordered lattices. the presented algorithm allows not only propagation of an existing crack but also initialization of new cracks. a0 a1 a3 times of initial procedures [s]: qc simplification 0.83 8.08 assemble stiffness matrix 1.09 5.81 0.15 assemble load vector 1.92 6.36 0.21 times of repeated procedures [s]: solve one step 3.99 0.25 0.15 update one step 1.08 2.31 0.12 table 2. time consumed by individual parts of simulation for various approaches. the presented example has shown that qc based approaches in combination with an adaptive algorithm are able to capture the exact macroscopic crack trajectory even if the area of crack propagation is not known at the beginning of simulation. the force-displacement diagram can be predicted with high accuracy in spite of the fact that the initial elastic response of simplified models is slightly stiffer. the adaptive extension of the area of high interest has turned out to be sufficient to obtain accurate results and a new triangulation of interpolation mesh is not necessary. the application of qc approaches provides a significant reduction of the number of unknown dofs and of links used in numerical models. this reduction leads to a significant speed-up of one simulation step, at the price of an increased cost of the initial preparation (model simplification) before the first step. if both interpolation and homogenization are applied, the speed-up factor of one step is 18.8 and the whole simulation of crack propagation (including 240 steps) gets more than 15 times faster. acknowledgements financial support received from the czech technical university (sgs project 17/043/ohk1/1t/11) and the czech science foundation (gačr projects 14-00420s and 1704150j) is gratefully acknowledged. references [1] e. b. tadmor, m. ortiz, r. phillips. quasicontinuum analysis of defects in solids. philosophical magazine a 73(6):1529–1563, 1996. [2] d. m. kochmann, j. s. amelang. the quasicontinuum method: theory and applications. in multiscale materials modeling for nanomechanics, chap. the quasicontinuum method: theory and applications, pp. 159–193. springer, 2016. [3] k. mikeš, m. jirásek. quasicontinuum method extended to irregular lattices. computers and structures 192:50–70, 2017. [4] r. e. miller, e. b. tadmor. a unified framework and performance benchmark of fourteen multiscale atomistic/continuum coupling methods. modelling and simulation in materials science and engineering 17(5):053001, 2009. 84 acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0026 acta polytechnica ctu proceedings 22:26–30, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app optimization of mix-design of chemically resistant sprayed concrete radek hermanna, ∗, rostislav drochytkab, jiří kolískoc, pavel reitermand a brno university of technology, admas centre, purkynova 139, 612 00 brno, czech republic b brno university of technology, faculty of civil engineering, veveri 95, 602 00 brno, czech republic c czech technical university in prague, klokner institute, solinova 7, 166 08 prague 6, czech republic d czech technical university in prague, faculty of civil engineering, thakurova 7, 166 29 prague 6, czech republic ∗ corresponding author: xherma23@vutbr.cz abstract. this paper addresses optimization of mix-design and granulometric curves of aggregate of chemically resistant cementitious sprayed mix. the target of the paper is to improve properties of currently developed mix-design, which has no chemical resistance, and to utilize the secondary raw materials. the mix was optimized by additions improving chemical resistance and binder substituted by secondary raw materials. the binder was substituted by finely ground waste glass and high-temperature fly ash. filler was substituted with slag from deposit yard and waste sand with soluble glass from steel manufacture. main observed properties for design of optimized mix were workability time, consistence and physical-mechanical properties. optimized mix-design successfully replaced all fine parts of filler and up to 30% of binder while the physical-mechanical properties were preserved or even enhanced. optimized mix-designs show the same consistencies and workability times. keywords: sprayed mix, optimization, secondary raw materials, water absorbing capacity, granulometric curve. 1. introduction the problem of sprayed mixes has been solved since the beginning of the 20th century, equipment for application of sprayed concrete was patented in 1911 under the name cement gun together with sprayed material, gunite. sprayed mixes are usually connected mainly with tunnel works, but use is not limited only to those applications. use of sprayed concrete has advantages mainly because of its high density, low porosity and overall high quality of sprayed material as such. this material is suitable mainly for applications, where visual properties of final surface are not required. design of sprayed mixes requires mainly optimization of granulometric curves of aggregate to ensure high density and at the same time very good pumpability of the mix. the important factor is sufficient amount of fines supported by higher proportion of binder and admixtures. sprayed mixes have their own test procedures, which are described in a set of standards čsn en 14488 [1]. current time focuses on reduction of ecological impacts caused by manufacture of building materials, for this reason, waste or secondary raw materials are used. slag. this solid, inorganic and partly melted residuum is produced during combustion of coal or combustible shales on fire grates and in furnaces. this material has high contents of mineral residua of original combustible substance. it is suitable for concrete and mixes; however, properties of slag are not homogeneous, which may considerably change its quality. slag very often contains unburned residua of coal, humins, free oxides of calcium and magnesium and other substances, which influences considerably its final behavior in synergy with cement composite. time of deposit is a very important factor; generally – the longer time of deposit, the higher quality of the slag as such. during deposit time, rain flushes out salts, which may have adverse effect on final mix and create for example efflorescence. another negative aspect of slag is the fact, many deposit yards are covered with naturally seeded plants, which grow through the deposited slag and final product contains high amount of organic combustible products and hence certain volume instability. in accordance with legislative requirements, slag shall contain the maximum of 10% of combustibles and for use in concrete, high content of amorphous sio2 is assumed [2]. sand with soluble glass. this material is a waste product from steel manufacture. it is a residual material from forms after form removing of cast iron products. this material consists mainly of very pure siliceous sand with content of 80–85% sio2, and lower amount of soluble glass. annual production of waste sand with soluble glass in the czech republic is over 30 000 tons. there is a generally prevailing opinion, that sand with soluble glass is not suitable for concrete and mortars because of the soluble glass, as this increases amount of alkalis in matrix and may 26 http://dx.doi.org/10.14311/app.2019.22.0026 http://ojs.cvut.cz/ojs/index.php/app vol. 22/2019 optimization of mix-design of chemically resistant sprayed concrete contribute to alkali-silica reaction (asr). however, in lower amounts it has no impact on matrix and is almost required for sprayed concrete because soluble glass as addition accelerates setting and hardening of concrete. specific weight is similar as that of pure silica in the interval of 2500 − 2650 kg·m−3. sand with soluble glass is suitable because of its granulometry, when used as screened by screens 0.5 mm. however, because of soluble glass, it forms agglomerates of size up to 16 mm without any external influence [3, 4]. glass. glass is a transparent solid material with high content of amorphous sio2, produced by manufacture from glass batch. the most common glass products are packing material and sheet glass, which is nowadays manufactured with flow method, when the glass is poured on a layer of melted tin and an endless sheet of glass is produced. recycled glass used as a secondary raw material for manufacture of concrete comes mainly from packing glass. however, glass from automotive industry is used too, in particular windscreens of automobiles, where it is necessary to separate flexible membranes between layers of glass forming the windscreen composite. waste glass produced during manufacture and glass from sorted waste containers is used for manufacture of glass to reduce cost as temperature of melting broken glass is lower than temperature of melting glass batch. unfortunately, not all waste glass can be reused for glass manufacture and glass waste is a serious ecological problem. most frequent glass is soda-lime glass with 50% of silica sand, 10% na2co3 (soda), 12% caco3 (lime stone), 18% crushed splinters of waste glass and 4% of other substances. this glass represents up to 80% of overall glass waste. because of its composition and content of sio2 it shows high pozzolanic activity, which, together with cement, supports formation of csh gels in cement matrix. studies of keren zheng [5] state that if size of glass grain is lower than 75 µm, the material shows sufficient pozzolanic activity. use of ground glass and generally glass in concrete presents problem of alkali silica reaction (asr), which is a reaction, when reaction of alkalis (cement) and amorphous silica material (glass) creates in cement matrix new forms of larger volume and these can gradually lead to static failure of the whole structure. gerry lee proved in his paper, if the glass is ground to higher specific surface and medium size of particles is lower than 300 µm, asr does not occur [5–7]. high temperature fly ash. high temperature fly ash is incombustible residuum formed in thermal power plants during combustion of pulverized coal at temperatures between 1200 and 1700 °c. to separate fly ash from waste gases, power plants use mostly highly effective electro-static separators; fly ash can be easily used for other purposes. as a reaction to growing demand for fly ash in building industry, standard csn en 450 [8] was elaborated, which prescribes properties and criteria of suitability of fly ash for concrete and mortars. suitability of fly ash for concrete and mortars requires minimal content of 45% amorphous sio2 and maximal content of free cao, mgo and alkalis. high temperature fly ash presents nowadays up to 70% of all energetic by-products. thanks to its composition, high temperature fly ash is similar to natural pozzolans. these pozzolanic properties combined with cement help the formation of csh gels. specific weight of fly ash is between 2100 and 2600 kg·m−3. fly ash influences mainly long-term physical-mechanical properties, which brings considerable enhancement and help enclose the structure cement matrix. this phenomenon also enhances chemical resistance in time [3]. 2. materials and methods for optimization, one of currently developed sprayed fine-grained mixes was selected, which did not show increased chemical resistance before adjustments. manufacturer supplies partial components; binder was cement cem i 42.5 r, additions, fillers, polypropylene fibers and admixtures for adjustments of specific properties of the material. composition of reference is in following table 1. mix was adjusted with secondary raw materials. part of the binder was substituted with finely ground recycled glass from soda-lime glass (glass) and high temperature fly ash (hfa). filler was substituted with sand containing soluble glass (scsg) and slag from former thermal power plant and nearby deposit yard (ssl). substitution filler, glass with soluble glass and slag were pre-treated with jawed crusher and subsequently with sieve analyzed so that final granulometric curve corresponded with that of original fillers as much as possible. substitution binders were pretreated with laboratory ball grinder so that specific surface was comparable with specific surface of cement, in particular 400 m2·kg−1(± 50 m2·kg−1). grain size of individual substitution binder is stated in following tab. 2. all tests and experiments were carried out in accordance with regulations and standards. mix-designs. in total, five reference materials were made with all content of binder and filler, the only modified part was content of additions influencing specific properties. out of these basic mix-designs, one representative was selected based on physicalmechanical properties after 28 days, in particular tensile strength, compressive strength and water absorbing capacity. this mix-design was then modified with secondary raw materials. the influence of these secondary raw materials on the physical-mechanical properties of the high strength cementitious composites was studied in cooperation of brno university of technology and klokner institute czech technical university in prague. as regards filler, 100% substitution of given size fraction was tested; as for binder, three levels of substitution were tested for each secondary raw material. after evaluation of physical-mechanical properties after 28 days, above mentioned binders and 27 r. hermann, r. drochytka, j. kolísko, p. reiterman acta polytechnica ctu proceedings component proportion in the mix [wt.%] portland cement cem i 42,5 r 17-25 silica sand 0-0.5 mm 13-18 silica sand 0-5.4 mm 38-45 waste from ground feldspar 4-9 ground lime stone 10-17 mix of additions 1-1.9 table 1. composition of reference mix. secondary raw materials < 0.01 mm 0.01 − 0.063 mm > 0.063 mm high temperature fly ash 28.84 63.35 7.87 recycled glass 30.71 53.12 16.18 table 2. grain size of substitution binders [wt.%]. fillers were combined in various ways to ensure maximal utilization of secondary raw materials. materials were mixed separately for each mix from individual raw materials and then mixed with mixing water in the proportion 0.14 kg of water per 1 kg of dry mix. amount of mixing water was slightly adjusted depending on given secondary raw materials to keep the same consistency in fresh state (tested on flow table in accordance with csn en 1015-3 [9]). physicalmechanical properties were tested on test specimens with dimensions 160 × 40 × 40 mm, which were placed in water after form removal; tests were carried out after 2, 7 and 28 days of curing. laboratory tests. fresh mixes were subjected to tests to determine density, consistency and time of setting and hardening. hardened specimens were subjected to tests of physical-mechanical properties – density, tensile strength, compressive strength and after 28 days also water absorbing capacity. 3. results and discussion physical-mechanical properties. fig 1 shows compressive strength of reference mix-design and other mix-designs with adjusted composition of additions. these adjusted mix-designs are labeled bet 0-3 and have completely identical composition and watercement ratio, the only change is amount of additions influencing specific properties of material. based on the results, one mix-design with the highest compressive strength and lowest water absorbing capacity at the same time will be selected, as the best chemical resistance of such mix-design can be presumed. selection is assumed after 28 days of maturation and selected mix-design will be subsequently optimized with secondary raw materials. the results imply that because of using additions affecting specific properties of the material, considerably higher compressive strengths were achieved only after two days while water-cement ratio was the same. to verify properties and determine exact results, suitable mix-design was selected after 28 days of maturation. above mentioned mix-designs were subjected to determination of water absorbing capacity. with few exceptions it was confirmed that the higher was the compressive strength value, the lower was water absorbing capacity after 28 days. mix-design bet 1 was selected for further optimization, as it showed the best results of physical-mechanical properties. this mix-design was then optimized with secondary raw material to achieve better results than material bet ref. following fig 2 shows compressive strengths of optimized mix-designs based on bet-1, from which, 5 out of 21 mix-designs were selected for further testing. compressive strengths of optimized mix-designs are pictured in contrast with reference mix-design bet ref. in optimized mix-designs, 100% of fine aggregate is replaced with waste sand with soluble glass from manufacture of steel (scsg) and subsequently, combinations were tested with waste glass (glass), high temperature fly ash (hfa) in the amount of 2030% of binder substitution (cement) and slag from deposit yard (ssl) as 100% replacement of coarse aggregate. test results show that when secondary raw materials are combined with additions, it is possible to achieve strengths which are comparable or even higher than those of reference mix after 28 days – combination of waste glass, high temperature fly ash and 30% substitution of cement. in this mix-design, pozzolanic properties of given materials proved. the results show that 20% substitution of cement with waste glass increased compressive strength by 4%, 20% substitution of cement with high temperature fly ash increased compressive strength by 7% and 30% substitution of cement with high temperature fly ash increased compressive strength by approx. 1% after 28 days. when slag from deposit yard was used, it was necessary to add more than 10% mixing water to keep the value of workability, which caused considerable reduction of compressive strength and volume weight, results are stated only for information. 28 vol. 22/2019 optimization of mix-design of chemically resistant sprayed concrete figure 1. compressive strengths of reference mix-design. figure 2. compressive strengths of optimized mix-designs. labeling of mix-design water absorbing capacity after 28 days [%] bet ref 10.44 bet x 10.29 bet 0 10.12 bet 1 9.37 bet 2 10.39 scsg 100 + glass 20 10.83 scsg 100 + glass 30 10.87 scsg 100 + hfa 20 10.85 scsg 100 + hfa 30 10.92 ssl 100 + scsg 100 + hfa 40 12.65 table 3. water absorption of individual mix-designs after 28 days. 29 r. hermann, r. drochytka, j. kolísko, p. reiterman acta polytechnica ctu proceedings following tab 3 show values of water absorbing capacity of samples made per individual mix-designs. water absorbing capacity was determined on fractions of test specimens after the compressive strength test. tab 3 shows water absorbing capacities of specimens made per individual mix-designs, compared to specimens made per reference mix-design. it is obvious at first sight that water absorbing capacity of optimized mix-designs is higher by 0.4 – 0.5 than reference sample. water absorbing capacity was determined on samples only 28 days old, when it is not expected that properties of used additions and pozzolanic properties of secondary raw materials show their effect. however, because of character of used raw materials it can be assumed, that the value of water absorbing capacity will decrease in time. 4. conclusion optimization of additions used in manufacture of above-mentioned sprayed mix could preserve or enhance physical-mechanical properties of reference material. reference material reached relatively high compressive strengths even after 2 days of curing; these strengths remained sufficient for sprayed application after optimization with secondary raw materials. use of secondary raw materials verified the hypothesis, that water absorbing capacity of matrix will be increased to certain extent. the usage of the selected secondary raw materials did not cause delay of setting or hardening. it was proved, that when 100% of finer aggregate was replaced with waste sand with soluble glass, combined with 30% replacement of binder (cement), the resulting mix shows after 28 days comparable or even better physical-mechanical properties and thanks to selection of secondary raw materials with pozzolanic properties it can be assumed, properties of the material will improve in the course of time. applicability of the material in more demanding exposition conditions is assumed after at least 60 days, when properties of additions and pozzolanic properties of used materials will manifest. because of pozzolanic additions, long-term properties are determined too, after 90 days. however, these are not yet available at the time of publishing this paper. selected mix-designs will be used for a test spraying and samples made in this way will be subjected to testing of physical-mechanical properties after 2, 7, 28 and 90 days including testing water permeability, resistance to chemical de-icing substances and resistance to chemical substances. the target is to create fine-grained sprayed mix with increased resistance to chemical substances and increased life time in demanding exposition conditions. acknowledgements this paper has been worked out under the project no. lo1408 “admas up – advanced materials, structures and technologies”, and under the project fast-s-18-5406 “study of structure of advanced silicate composites focusing on characteristics of action of secondary raw materials”. references [1] čsn en 14488-6: testing sprayed concrete part 6: thickness of concrete on a substrate. standard, czech office for standards, metrology and testing, 2007. in czech. [2] shotcrete design guidelines: design guidelines to avoid, minimise and improve the appearance of shotcrete. tech. rep., roads and traffic authority of nsw, 2005. [3] p. pytlík. technologie betonu. vutium, brno, 2nd edn., 2000. [4] l. svoboda, et al. technologie betonu. luboš svoboda, praha, 3rd edn., 2013. [5] c. shi, k. zheng. a review on the use of waste glasses in the production of cement and concrete. resources conservation and recycling 52:234–247, 2007. doi:10.1016/j.resconrec.2007.01.013. [6] v. bílek. možnosti využití odpadních písků z výroby vodního skla ve stavebnictví. master’s thesis, brno university of technology, faculty of chemistry, 2013. [7] h. lee, a. hanif, m. usman, et al. performance evaluation of concrete incorporating glass powder and glass sludge wastes as supplementary cementing material. journal of cleaner production 170:683–693, 2018. doi:10.1016/j.jclepro.2017.09.133. [8] čsn en 450-1: fly ash for concrete part 1: definition, specifications and conformity criteria. standard, czech office for standards, metrology and testing, 2005. in czech. [9] čsn en 1015-3: methods of test for mortar for masonry part 3: determination of consistence of fresh mortar (by flow table). standard, czech office for standards, metrology and testing, 2000. in czech. 30 http://dx.doi.org/10.1016/j.resconrec.2007.01.013 http://dx.doi.org/10.1016/j.jclepro.2017.09.133 acta polytechnica ctu proceedings 22:26–30, 2019 1 introduction 2 materials and methods 3 results and discussion 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0108 acta polytechnica ctu proceedings 12:108–111, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app interactive motion platforms and virtual reality for vehicle simulators evžen thöndel pragolet, s.r.o., director and czech technical university in prague, assistant professor at the department of electric drives and traction correspondence: thoendel.jr@pragolet.cz abstract. interactive motion platforms are intended for vehicle simulators, where the direct interaction of the human body is used for controlling the simulated vehicle (e.g. bicycle, motorbike or other sports vehicles). the second use of interactive motion platforms is for entertainment purposes or fitness. the development of interactive motion platforms reacts to recent calls in the simulation industry to provide a device, which further enhances the virtual reality experience, especially with connection to the new and very fast growing business in virtual reality glasses. the paper looks at the design and control of an interactive motion platform with two degrees of freedom to be used in virtual reality applications. the paper provides the description of the control methods and new problems related to the virtual reality sickness are discussed here. keywords: motion platform, motion cueing, simulator, simulation, virtual reality, vr sickness. 1. introduction interactive motion platforms are intended for a simulation where direct interaction with a human body is required. this type of simulators is e.g. a motorbike or bicycle simulator or other types of vehicles where direct interaction with the human body is used for controlling the vehicle. a simulator with an interactive motion platform is a closed loop system, where a human being is a part of the system. in contrast to standard humanin-the-loop systems[1], the human does not interact with the system through control elements (such as a steering wheel or pedals) but creates forces and torques which are directly measured and used as input to a mathematical model. the structure of the closed loop system is shown in the next picture. the structure and principle of individual blocks are further discussed. 2. motion platform the purpose of a motion platform in simulators is to reproduce motion effects, such as acceleration, and position towards gravity. different types and principles are discussed in the literature[2]. the important parameter of each motion platform is the number of degrees of freedom (dof) and the range of motion in all directions. the most widely used motion platforms today have 2, 3 or 6 dof. in the case of simulators of land vehicles, the dominant motion is rotation around the horizontal axis. this motion can also be used for reproducing linear accelerations (by using a motion cueing algorithm[3]). the second aspect which must be considered, when designing a new motion platform, is the final price which significantly grows with the number of degrees of freedom. a motion platform with two degrees of freedom is the first design and implementation of an interactive motion platform. an extension with a horizontal rotation of 360 degrees is considered in the future, to avoid or to solve a problem connected with virtual reality sickness. the design of the 2dof interactive motion platform is shown in the picture below. figure 1. design of the motion platform. all kinematic and dynamic parameters are summarized in the following table: maximum payload 150 kg tilt angle 16 degrees degrees of freedom 2 (pitch and roll) maximum tilt rate 100 degrees/s maximum tilt acceleration 500 degrees/s2 power supply requirements 230 v/50 hz, 16 a weight 80 kg table 1. kinematic and dynamic properties. 108 http://dx.doi.org/10.14311/app.2017.12.0108 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 interactive motion platforms and virtual reality for vehicle simulators figure 2. structure of the system with an interactive motion platform. 3. mathematical model of the dynamic response a mathematical model of the dynamic response reads information about the acting forces and torques and calculates the response of the motion platform. besides the input from the human, the mathematical model of the dynamic response can also read signals from a superior mathematical model where the dynamic response of the simulated vehicle is implemented. an implementation of the mathematical model of the dynamic response is shown in the next picture. this structure applies a motion cueing algorithm as well as a model of direct dynamic response (called stiffness in the following structure). the principle of a motion cueing algorithm is described in detail in the previous author’s publications[4], and will not be discussed here. a model of the direct dynamic response can be any dynamic system which reacts directly to forces measured on the motion platform. one possible use of the interactive motion platform is rehabilitation or fitness where the model can implement a system of the load. one example is presented in the next picture where a simple model of a rubber ball is shown. all mathematical models have been implemented in matlab/simulink and directly compiled to the control hardware. this method of programming is very effective, and it allows the motion platform to also be used in research projects[5]. 4. load sensors as discussed above, the main innovation of the new motion platform is the possibility to measure forces and torques which are created by a person standing or sitting on the motion platform. four load cells, weighing 100 kg each, are used to measure and further calculate the position (center of gravity) and the intensity of the loading force. figure 3. load sensors attachment. 5. the use in virtual reality there has been fast growth in virtual reality over the last two years, especially connected to the use of virtual glasses. new applications and technologies are continuously being researched. the presented motion platform is ideally suited for vehicle simulators where virtual glasses are used to generate visual perception. several designs of applications have been created, and the first real application of a kayak simulator has been finished in 2016[6]. the first test and the response from many users show a promising potential for business success and further technological enhancement of the use of vr glasses because it significantly reduces the occurrence of virtual reality sickness. 6. virtual reality sickness virtual reality sickness is a common problem which is often discussed. any mismatch between what is seen and what is felt can lead to virtual reality sickness. the motion platform can reproduce a wide range of motion effects, and therefore it contributes to the reduction of the probability of occurrence of virtual reality sickness. however, two degrees of freedom do not cover all motion directions. the horizontal rotation especially is often mentioned in relation to this phenomenon, and it is recommended not to rotate the camera in the vr world at all[7] or to use a 109 evžen thöndel acta polytechnica ctu proceedings figure 4. mathematical model of the dynamic response. figure 5. mathematical model of the direct dynamic response – rubber ball. figure 6. application of the motion platform. 110 vol. 12/2017 interactive motion platforms and virtual reality for vehicle simulators motion platform with a rotary axis[8]. the vertical linear acceleration can partially be compensated for by rotating the motion platform. therefore, it is planned to extend the construction with a rotation of 360 degrees. this new design will be an optimal construction for simulators using vr headsets with respect to the range of reproduction of motion effects and the financial costs. 7. conclusion a new interactive motion platform has been presented here. the possibility to measure acting forces and torques makes the motion platform suitable for the use in vehicle simulators where direct interaction with a human body is required (bicycle or motorbike). the main advantage is the enhancement of the experience and the contribution to the reduction of the virtual sickness phenomenon. we plan to extend the motion platform with a rotary axis which further improves the solution of the problem with vr sickness. acknowledgements the research and development of the new interactive motion platform is supported by the company pragolet, s.r.o. member of the center of intelligent drives and advanced machine control (cidam) supported by the technology agency of the czech republic (tacr). references [1] human-in-the-loop. wikipedia, the free encyclopedia. wikipedia, the free encyclopedia, 10 dec. 2016. web. 10 dec. 2016. [2] e. thöndel. design and optimal control of a linear electromechanical actuator for motion platforms with six degrees of freedom. 2011, intelligent automation and systems engineering, springer, 65-77. [3] e. thöndel. design and optimization of a motion cueing algorithm for a truck simulator. 2012, in proceeding of the 2012 european simulation and modeling conference, essen, germany, 165-170. [4] e. thöndel. model predictive motion cueing algorithm for a truck simulator. 2013, in proceeding of the 2013 european simulation and modeling conference, lancaster, united kingdom, 188-192. [5] p. members. vývoj zpětnovazebního rehabilitačního přístroje založenãľho na principu pohybové terapie. research project, https://www.rvvi.cz/cep?ss=detail&h=ta03010313, 2015. [6] p. contributors. 4d vr experience modul. the global association for marketing at retail, popai awards 2016, http://www.popai.cz/files/tinymce/files/ popai%20awards%202016/vyher/20.pdf, 2016. [7] o. contributors. oculus rift vr motion sickness. oculus http: //riftinfo.com/oculus-rift-motion-sickness/ -11-techniques-to-prevent-it, 2016. [8] r. contributors. yaw motion sickness. http://www.rotovr.com, 2016. 111 https://www.rvvi.cz/cep?ss=detail&h=ta03010313 http://www.popai.cz/files/tinymce/files/popai%20awards%202016/vyher/20.pdf http://www.popai.cz/files/tinymce/files/popai%20awards%202016/vyher/20.pdf http://riftinfo.com/oculus-rift-motion-sickness/-11-techniques-to-prevent-it http://riftinfo.com/oculus-rift-motion-sickness/-11-techniques-to-prevent-it http://riftinfo.com/oculus-rift-motion-sickness/-11-techniques-to-prevent-it http://www.rotovr.com acta polytechnica ctu proceedings 12:108–111, 2017 1 introduction 2 motion platform 3 mathematical model of the dynamic response 4 load sensors 5 the use in virtual reality 6 virtual reality sickness 7 conclusion acknowledgements references student conference on nuclear engineering šimáně 2017 6th 7th november, 2017, prague, czech republic acta polytechnica ctu proceedings organised by: department of nuclear reactors, faculty of nuclear science and engineering, ctu in prague simane.fjfi.cvut.cz editors: martin ševeček martin cesnek scientific comittee: milan štefánik martin ševeček martin cesnek local organizing comittee: adam kecek vojtěch caha martin cesnek filip fejt martin ševeček milan štefánik guarantor of the peer review process: milan štefánik guarantor of language editing: martin ševeček iv vol. 14/2018 preface nuclear science and technology have a long history within the czechoslovakia. even though the czech republic and slovakia have been separated since 1993, wide cooperation on different levels including education and research still exists. in order to support cooperation and good relations between czech and slovak students, the organizing committee decided to hold the 2nd student conference on nuclear engineering šimáně 2017. the conference is named after professor šimáně, who graduated at the dr. edvard beneš technical university in brno. during his early career, he had the opportunity to work with fréderic joliot-curie at college de france. later, he significantly contributed to the foundation of the nuclear science, industry and education in the czechoslovakia. he became the first employee of the institute of nuclear physics at the czechoslovakian academy of science, later he became the first director of the nuclear research institute in řež. professor šimáně was well established even on the international level. he was a director at the joint institute of nuclear research in dubna or a division director at the international agency for atomic energy. his work was also significantly connected with our faculty where he acted as dean from 1967 to 1972. the main objective of the conference is to provide an opportunity for bsc., msc. and ph.d. students of nuclear-oriented study programmes to publish and present their scientific results achieved during their university studies. furthermore, it should give the participants experience in the field of presenting their own research, writing scientific papers and also broaden their knowledge throughout the related fields. the organizing committee would like to thank all participants for their presentations and also to the ctu in prague for funding this student conference. the committee also hopes that the work achieved and presented will fulfill the heritage of professor čestmír šimáně. v acta polytechnica ctu proceedings list of participants: dana barátová stu fei martin lovecký zču dana_baratova@stuba.sk lovecky@rice.zcu.cz pavel bartl čvut fjfi/kjch tomáš matlocha čvut fjfi pavel.bartl@fjfi.cvut.cz matlocha@ujf.cas.cz ivana bonková uk ba ondřej novák čvut fjfi ivana.bonkova@gmail.com ondrej.novak2@fjfi.cvut.cz viktor burian uk mff adela nováková čvut fjfi viktor.burian@protonmail.com novakad3@fjfi.cvut.cz vojtěch caha čvut fjfi sebastian nývlt čvut fjfi vojtacaha@gmail.com nyvltseb@fjfi.cvut.cz daniel campolina čvut fjfi filip osuský stu fei campolina@cdtn.br filip.osusky@stuba.sk lucie celbová čvut fjfi/kipl pavel suk čvut fjfi celboluc@fjfi.cvut.cz sukpave2@fjfi.cvut.cz martin cesnek čvut fjfi jakub svoboda čvut fjfi martin.cesnek@fjfi.cvut.cz svoboj68@fjfi.cvut.cz tomáš czakoj čvut fjfi martin ševeček čvut fjfi tomas.czakoj@gmail.com martin.sevecek@fjfi.cvut.cz michal farník čvut fjfi/kf václav šísl čvut fjfi farnimi1@fjfi.cvut.cz sislvacl@fjfi.cvut.cz filip fejt čvut fjfi/kjr jan škarohlíd čvut fs filip.fejt@fjfi.cvut.cz jan.skarohlid@fs.cvut.cz markéta floriánová čvut fjfi/kjch jan vimpel čvut fs jan.vimpel@seznam.cz dominika holková stu fei daniel vlček čvut fjfi dominika.holkova@stuba.sk daniel.vlcek1@gmail.com kateřina jiráková čvut fjfi lenka volfová uk přf jirakkat@fjfi.cvut.cz volfl@fzu.cz adam kecek čvut fjfi/kjr pavel vrbka čvut fjfi adam.kecek@fjfi.cvut.cz pavel.vr@seznam.cz jakub krejčí čvut fjfi jiří závorka čvut fs krejci@ujp.cz jiri.zavorka@fs.cvut.cz denisa kubániová uk mff kubaniova@mbox.troja.mff.cuni.cz vi acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0109 acta polytechnica ctu proceedings 13:109–114, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app variationally-based effective dynamic thickness for laminated glass beams jaroslav schmidt∗, alena zemanová, tomáš janda, jan zeman, michal šejnoha czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jarasit@gmail.com abstract. laminated glass, consisting of glass layers connected with transparent foils, has found its applications in civil, automotive, or marine engineering. due to a high contrast in layer properties, mechanical response of laminated glass structures cannot be predicted using classical laminate theories. on the other hand, engineering applications demand easy-to-use formulas of acceptable accuracy. this contribution addresses such simplified models for free vibrations of laminated glass beams, with the goal to determine their natural frequencies and modal damping properties. our strategy is to approximate the complex behavior of a laminated structure with that of an equivalent monolithic beam. its effective thickness is determined by the variational method proposed by galuppi and royer-carfagni for static problems, which we extended for modal analysis. we show that this new approach overcomes inaccuracies of the currently used dynamic effective thickness model by lópez-aenlle and pelayo. keywords: laminated glass, dynamic effective thickness, modal analysis. 1. motivation in the past, glass was seen as an infill material for transparent areas with no load-bearing function in a structural system. the invention of laminated glass became one of the milestones that allowed glass to attain a genuine structural role [1]. starting from windscreens in the car industry, its use has expanded into other fields of industry, including civil engineering. today, layered glass is used, for example, for floor and roof systems, columns, staircases etc. laminated glass is a layered composite material consisting of several layers of glass and polymer. apart from its primary function in enhancing the post-peak load-bearing capacity, the polymer interlayer also provides damping of vibrations [2]. therefore, it can also reduce the possible noise and vibration problems, which can occur in structures under dynamic loading. on the other hand, the nature of the polymer response is strongly frequency-dependent and also temperature-sensitive, e.g., [3], [4], or [5]. its behavior is often described by a complex-valued dynamic shear modulus, see figure 1. this complex-valued formulation allows us to transmit the information about the resonance behavior and damping of the structural element in a single complex number. for these reasons, the eigenvalue problem which describes the free vibrations of laminated glass becomes complex and nonlinear; see section 2.1 for more details. to the best of our knowledge, such type of problems cannot be currently addressed with standard finite element systems used in civil engineering. as engineering practice demands easy-to-use approaches of acceptable accuracy, this contribution addresses simplified models for free vibrations of laminated glass beams, with the goal to determine their natural frequencies and modal damping properties. our focus is on the effective thickness approaches – one group of the simplified methods popular in the design of laminated glass structures, e.g., [6], [7], [8]. in particular, we demonstrate on selected examples the (in)accuracy of the state-of-the-art effective thickness approaches from literature, and we discuss our proposal how to increase their accuracy by employing principles of variationally-based effective thickness methods. the structure of the paper is as follows. in section 2, we briefly describe all approaches used in this study. the examples are introduced and the results of our comparison are discussed in section 3. finally, the main findings of our study are summarized in section 4. 2. methods three methods for vibration analysis of laminated glass beams are used in this section. the first method is based on solving the nonlinear and complex eigenvalue problem using a finite element discretization and the newton method. this method provides us with a reference solution to assess the accuracy of effective thickness approaches. the second method is the dynamic effective thickness method (det) proposed by lópez-aenlle and pelayo [6]. finally, the last method is derived from the enhanced effective thickness approach (eet) by galuppi and royer-carfagni for deflections under static loading [7], which we adjusted for modal analysis. 109 http://dx.doi.org/10.14311/app.2017.13.0109 http://ojs.cvut.cz/ojs/index.php/app j. schmidt, a. zemanová, t. janda et al. acta polytechnica ctu proceedings 10−1 100 101 102 103 frequency [hz] 10−3 10−2 10−1 100 101 102 103 s h e a r m o d u l u s [m p a ] 25◦c 50◦c storage modulus loss modulus figure 1. the frequency dependence of the real and imaginary part of the complex shear modulus for polyvinyl butyral (pvb) according to [3]. for the sake of simplicity, we focus on the most common case of laminated glass configuration – a threelayered beam. however, the generalization of all methods for multi-layered beams can be also found in the literature [9]. in our analysis, we consider a threelayered beam with the same material parameters for both glass layers and with constant properties along the length according to figure 2. then, the crosssection areas ai and the second moments of area ii are defined by the standard relations ii = 1 12 bh3i , ai = bhi, i = {1, 2, 3}, (1) where b and hi stand for the width and the thicknesses of the layers; see figure 2. 2.1. reference method (rm) our reference method is based on the finite element method (fem) together with the newton method due to the nonlinearity of the solved eigenvalue problem. the system of equations of natural vibrations is written in the fem matrix form as follows( k(ω) −ω2m ) u = 0, (2) where ω is the natural angular frequency, u is the corresponding vector of the mode shape, k(ω) is the complex frequency-dependent stiffness matrix, and m is the real constant mass matrix. the eigenvalue problem described by eq. (2) specifies the mode shape up to an arbitrary constant. therefore, we use a regularization condition u 0 t(u − u 0) = 0, (3) in order to ensure the uniqueness of the solution. here, u 0 is the initial mode shape solving the real eigenvalue problem with the constant stiffness matrix accounting only for the initial shear modulus of the polymer interlayer. for each initial eigen-pair ω0 and u 0, we solve iteratively a linearized system of equations, derived by the newton method [10] from eq. (2) and eq. (3), until the convergence is achieved. then, the natural frequencies fhz and the loss factors η (representing damping) correspond to the real part and to the ratio of imaginary and real part of the complex eigen-frequencies fhz = √ re[ω2] 2π , (4a) η = im[ω2] re[ω2] . (4b) this numerical solution converges for a small stopping tolerance to the exact solution. that is why this method is taken as the reference one, when assessing the accuracy of the simplified methods considered next. 2.2. dynamic effective thickness (det) to our best knowledge, only one effective thickness approach for modal analysis of laminated glass beams can be found in literature. this approach by lópezaenlle and pelayo [6] is derived from a closed-form solution of free vibration for a three-layer beam with simply-supported ends. for the other boundary conditions, the authors adjust the formulation using the wavenumbers of an euler-bernoulli beam, see ahead eq. (7) and table 2. the analytical expression for this dynamic effective thickness has the following form hef(ω) = 3 √√√√(h31 + h33) ( 1 + y 1 + h1 q(ω)(h1+h3) ) , (5) where y is a geometric parameter, which depends only on the thicknesses of layers and is given by y = 12h1h3(0.5h1 + h2 + 0.5h3) (h31 + h33)(h1 + h3) , (6) and q(ω) is a material parameter given by the identity q(ω) = g2(ω) e1h3h2β2n . (7) in this equation, g2(ω) is the frequency-dependent complex shear modulus of the interlayer, e1 is the young modulus of glass layers, and βn is the wavenumber corresponding to the given boundary conditions and to the n-th mode, see ahead table 2. because the shear modulus g2(ω) is complex-valued, the effective thickness from eq. (5) is also complex. then, the complex-valued natural angular frequencies can be expressed analytically, similarly as for the monolithic beam ω2 = β4ne1h 3 ef(ω) 12m̄ , (8) 110 vol. 13/2017 variationally-based effective dynamic thickness h1 h2 h3 l b figure 2. the layout of a three-layered laminated glass beam. where m̄ = ρ1h1 + ρ2h2 + ρ3h3 (9) is the effective mass per unit of length and width. the natural frequencies fhz and loss factors η are obtained from this complex-valued angular frequency ω according to eq. (4). 2.3. enhanced effective thickness (eet) firstly, we introduce the formula for the deflectioneffective thickness according to galuppi and royercarfagni [7], and subsequently, we propose its extension to the free vibration analysis. to shorten the notation, h is defined as a distance of centerlines of glass layers, so that h = h2 + 0.5(h1 + h3). (10) further, itot is the second moment of area of glass layers for the monolithic case itot = i1 + i3 + a1a3 a1 + a3 h2, (11) and is is the last term from (11) divided by the width of the beam b, thus is = h1h3 h1 + h3 h2. (12) using these quantities, the enhanced effective thickness for the deflection under static loading from [7] has the following form hef = 3 √√√√√ 1ζ h31 + h33 + 12is + 1 − ζ h31 + h33 , (13) where ζ is the coefficient of shear cohesion ζ = 1 1 + i1 + i3 µitot a1a3 a1 + a3 ψ , (14) and µ is the non-dimensional ratio of the glass and interlayer stiffnesses µ = g2b e1h2 . (15) finally, ψ is a coefficient which depends on the shape of the deflection curve and is described by the following equation ψ = ∫ l 0 (g ′′(x))2dx∫ l 0 (g ′(x))2dx , (16) geometry length l = 1 m width b = 0.1 m thicknesses of glass h1 = h3 = 10 mm thickness of interlayer h2 = 1.52 mm glass density ρ1 = ρ3 = 2500 kg/m3 poisson’s ratio ν1 = ν3 = 0.22 young’s modulus e1 = e3 = 72 gpa interlayer density ρ2 = 1100 kg/m3 poisson’s ratio ν2 = 0.49 prony series for gmm taken from [3] table 1. the properties of laminated glass beams (gmm – generalized maxwell model). where the deflection curve is represented by the function g(x), which describes the shape of the curve regardless of the stiffness of the beam. this deflection-effective thickness was derived by minimization of the strain energy functional for a laminated glass beam in bending under static loading. for its use in modal analysis, we made just two intuitive adjustments of this method: (1.) we used the complex-valued interlayer shear modulus g2(ω) in eq. (15). (2.) we replaced the shape function of the deflection under static loading in eq. (16) with the one corresponding to the n-th mode shape of a monolithic beam under given boundary conditions. this adjustment leads to a complex effective thickness in eq. (13), similarly to the det approach. for the evaluation of the natural frequencies and the damping, we again use eq. (4). 3. comparison of methods 3.1. examples in this section, we discuss the applicability of effective thickness approaches in modal analysis and how our adjustment changes the errors of the method. the comparison is provided for three types of boundary conditions: simply-supported, fixed-fixed, and free-free. 111 j. schmidt, a. zemanová, t. janda et al. acta polytechnica ctu proceedings mode shape functions simply-supported beam g(x) = sin nπx l fixed-fixed beam g(x) = cosh βnx− cos βnx− sinh βnl + sin βnl cosh βnl− cos βnl (sinh βnx− sin βnx) free-free beam g(x) = sinh βnx + sin βnx− sinh βnl− sin βnl cosh βnl− cos βnl (cosh βnx + cos βnx) table 2. the overview of the n-th mode shape functions for the coordinate x ∈ 〈0, l〉 for a simply-supported, fixed-fixed, and free-free beam. (the products of the wavenumbers and the beam length βnl for the first three modes are β1l = 4.7300, β2l = 7.8532, and β3l = 10.996.) the geometry and the material properties of glass and interlayer appear in table 1. whereas glass is treated as an elastic material, the viscoelastic behavior of polymer, which is in our case polyvinyl butyral (pvb), is described using a generalized maxwell model, whose parameters are taken from [3]. due to the temperature-sensitiveness of pvb, two different ambient temperatures were assumed – the room temperature 25◦c and the temperature 50◦c corresponding to an external panel under summer sunlight. 3.2. evaluation of shape coefficients recall that our extension of the enhanced effective thickness approach [7] to the modal analysis of laminated glass beams requires the evaluation of the shape coefficients ψ according eq. (16). for these coefficients, we need mode shape functions corresponding to the three given boundary conditions. their overview taken from [11] is shown in table 2. for a simplysupported beam, one closed-form expression can be written for all mode shape functions; however, the formulas for the other boundary conditions are more involved and contain a product of the wavenumbers βn and the beam length l that needs to be evaluated numerically from a characteristic equation for free vibrations [11]. then, we used the integral formula eq. (16) and evaluated, for all boundary conditions, the shape coefficients ψ summarized in table 3 using the mode shape functions for a monolithic euler-bernoulli beam from table 2. with these results at hand, we can calculate the complex enhanced effective thickness and subsequently the natural frequencies and the loss factors for our comparison. 3.3. results and discussion in this section, the results obtained by the reference method (rm), the effective thickness approaches from [6] (det) or from [7] (eet), their errors, and their comparison are shown. the first three natural frequencies fhz and loss factors η are summarized in table 4. the errors of the effective thickness approaches are evaluated against the reference method based on the finite element complex-eigenvalue solver. shape coefficients ψ beam mode 1 2 3 simply-supported π2 l2 (2π)2 l2 (3π)2 l2 fixed-fixed 40.7 l2 82.6 l2 148 l2 free-free 10.1 l2 34.9 l2 78.2 l2 table 3. summary of the shape coefficients ψ for three different boundary conditions and the first three mode shapes. it is evident from the first two subtables in table 4 that the adjusted eet approach gives the same results as the det method for the simply-supported beam for both temperatures. both effective thickness methods give the errors in natural frequencies less than 1% and in loss factors less than 10%. for the room temperature, even the errors in loss factors are under 1%. for the two other boundary condition, the det approach provides quite good predictions for natural frequencies with the errors under 13%. however, the errors in loss factors can be very high (up to 80%). therefore, this method is mostly unable to provide a good estimation of damping. the adjusted eet approach helps to overcome these inaccuracies. for the fixed-fixed beam, this approach gives the largest error in natural frequencies 3% (against the 8% error by the det) and in loss factors 9% (against the 44% error by the det). for the free-free beam, the largest value of errors in natural frequencies is 3% (against the 13% error by the det) and in loss factors 19% (against the 78% error by the det). this demonstrates that our intuitive extension of the variationally-based enhanced effective thickness concept from [7] provides better (or for the simplysupported beam the same) estimates of both the natural frequencies and loss factors for laminated glass beams. 112 vol. 13/2017 variationally-based effective dynamic thickness simply-supported beam at 25◦c mode natural frequency [hz] error [%] mode loss factor [%] error [%] rm det eet det eet rm det eet det eet 1 52.03 52.07 52.07 0.09 0.09 1 1.93 1.93 1.93 -0.25 -0.25 2 197.9 198.5 198.5 0.30 0.30 2 4.24 4.23 4.23 -0.32 -0.32 3 419.1 421.4 421.4 0.56 0.56 3 6.15 6.14 6.14 -0.20 -0.20 simply-supported beam at 50◦c mode natural frequency [hz] error [%] mode loss factor [%] error [%] rm det eet det eet rm det eet det eet 1 30.81 30.97 30.97 0.53 0.53 1 24.41 22.17 22.17 -9.18 -9.18 2 108.1 108.5 108.5 0.42 0.42 2 16.89 16.25 16.25 -3.81 -3.81 3 232.0 233.1 233.1 0.47 0.47 3 14.46 13.90 13.90 -3.84 -3.84 fixed-fixed beam at 25◦c mode natural frequency [hz] error [%] mode loss factor [%] error [%] rm det eet det eet rm det eet det eet 1 110.2 115.4 111.3 4.78 1.09 1 5.42 3.09 5.08 -43.05 -6.27 2 285.3 301.4 292.7 5.65 2.59 2 7.15 5.29 6.50 -25.94 -9.02 3 527.2 556.3 541.1 5.53 2.65 3 7.86 6.74 7.56 -14.28 -3.82 fixed-fixed beam at 50◦c mode natural frequency [hz] error [%] mode loss factor [%] error [%] rm det eet det eet rm det eet det eet 1 60.06 64.63 60.21 7.62 0.25 1 12.99 18.74 12.72 44.19 -2.14 2 160.2 164.6 160.9 2.79 0.46 2 11.51 14.91 11.80 29.52 2.49 3 308.8 314.0 310.2 1.66 0.43 3 10.98 12.89 10.87 17.39 -1.03 free-free beam at 25◦c mode natural frequency [hz] error [%] mode loss factor [%] error [%] rm det eet det eet rm det eet det eet 1 117.8 115.4 118.4 -2.00 0.56 1 1.73 3.09 1.50 78.05 -13.64 2 307.9 301.4 314.1 -2.12 2.00 2 4.13 5.29 3.37 28.20 -18.41 3 568.4 556.3 584.5 -2.12 2.84 3 5.77 6.74 5.04 16.74 -12.64 free-free beam at 50◦c mode natural frequency [hz] error [%] mode loss factor [%] error [%] rm det eet det eet rm det eet det eet 1 73.95 64.63 74.62 -12.60 0.90 1 26.44 18.74 25.39 -29.13 -3.96 2 170.5 164.6 175.6 -3.45 2.97 2 20.85 14.91 22.20 -28.51 6.48 3 322.5 314.0 325.3 -2.66 0.86 3 17.57 12.89 18.20 -26.61 3.62 table 4. natural frequencies and loss factors for the three-layer laminated glass beam for the first three modes determined by the reference method (rm), the dynamic effective thickness method (det), and the adjusted enhanced effective thickness approach (eet) with their errors against the reference method (rm). 113 j. schmidt, a. zemanová, t. janda et al. acta polytechnica ctu proceedings 4. conclusions finally, we would like to summarize that in this paper • we verified the effective thickness approaches against the finite element complex-eigenvalue solver, and • we proposed a new dynamic effective thickness concept derived from [7]. it follows from our comparison that this extended enhanced effective thickness method • provides the same results for the simply-supported beam as the dynamic effective thickness from [6] with the errors less than 1% in natural frequencies and 10% in loss factors, and • improves the estimates of natural frequencies and loss factors for the other two boundary conditions with the errors less than 3% in natural frequencies and 19% in loss factors. in our future work, we would like to justify this intuitive approach by rigorously extending the variationally-based derivation of the effective thickness from [7] to the dynamic problems. acknowledgements this publication was supported by the czech science foundation, the grant no. 16-14770s and by the grant agency of the czech technical university in prague, grant no. sgs17/043/ohk1/1t/11. references [1] m. haldimann, a. luible, m. overend. structural use of glass, vol. 10 of structural engineering documents. iabse, 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[11] r. w. clough, j. penzien. dynamics of structures. computers & structures, inc, 2003. 114 http://dx.doi.org/10.1016/j.compstruct.2016.05.105 http://dx.doi.org/10.1016/j.polymertesting.2014.04.011 http://dx.doi.org/10.1016/j.euromechsol.2017.01.006 http://dx.doi.org/10.1016/j.conbuildmat.2014.04.003 http://dx.doi.org/10.1016/j.engstruct.2011.12.039 http://dx.doi.org/10.1177/1099636217695479 acta polytechnica ctu proceedings 13:110–115, 2017 1 motivation 2 methods 2.1 reference method (rm) 2.2 dynamic effective thickness (det) 2.3 enhanced effective thickness (eet) 3 comparison of methods 3.1 examples 3.2 evaluation of shape coefficients 3.3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2016.6.0028 acta polytechnica ctu proceedings 6:28–33, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app mapping the surroundings as a requirement for autonomous driving martin steininger∗, christoph stephan, christian böhm, fabian sauer, roland zink faculty of electrical, media and computer engineering, deggendorf institute of technology, 94469 deggendorf, germany ∗ corresponding author: martin.steininger@stud.th-deg.de abstract. motivated by the hype around driverless cars and the challenges of the sensor integration and data processing, this paper presents a model for using a xbox one microsoft kinect stereo camera as sensor for mapping the surroundings. today, the recognition of the environment of the car is mostly done by a mix of sensors like lidar, radar and cameras. in the case of the outdoor delivery challenge robotour 2016 with model cars in scale 1 : 5, it is our goal to solve the task with one camera only. to this end, a three-stage approach was developed. the test results show that our approach can detect and locate objects at a range of up to eight meters in order to incorporate them as barriers in the navigation process. keywords: microsoft kinect, mapping, 3d mapping, robotour, autonomous driving. 1. introduction in addition to the extensive launch of full electric vehicle (fev) autonomous driving is deemed to be a key research and development trend in the automotive industry. the entire industry is working on the topic and has already developed first solutions for automatic braking, line keeping (i.e. audi “active lane assist” or bmw “active assist”) or even autopilots for driverless cars like tesla motors’ “autopilot” or google’s “driverless car” [1]. the used navigation and self-driving algorithms run on maps [2] and require very accurate digital (geo-)data of the environment [3]. therefore, the spatial coverage of the surroundings as a basis for autonomous navigation of the vehicle is common to all applications. consequently, autonomous driving – regardless of the respective task such as automatic braking, adaptive cruise control, lane keeping or autonomous driving – requires detection and imaging of the space which it will pass through. photo cameras, lidar-, radaror other sensors continuously detect and measure the surrounding of the car, creating a digital spatial image of the area (map). combined with positioning services like gps or galileo, the location of the vehicle can be determined and set in relation to surrounding objects such as other vehicles or obstacles. however, it is not an ordinary cartographic map resulting from the permanent mapping process but digital images that are stored temporarily as server frames for the navigation. the navigation of the car finally relies on mathematical functions (algorithms) to pass through the digital and real space on the shortest (spatial) or fastest (time) way. this paper addresses the issue of the space detection which is considered as an essential basis of autonomous driving and robot navigation (see for example [4]) and describes the use of a microsoft x-box one kinect stereo camera for autonomous driving for the outdoor delivery challenge robotour 2016 (robotika.cz). we present an approach for mapping the surroundings including the three steps (1) 3d capturing, (2) projection and filtering and (3) object detection. the hardware and the approach are implemented on a four wheel drive off-road buggy, scale 1 : 5. the main hardware consists of a congatec ts180 com express module. for the main processing, it communicates with the x-box one kinect and a raspberry pi and displays the crucial information on a krämer v-800 touchscreen. the raspberry pi reads the sensor values from the gps and the bosch bno055 and transmits them to the congatec module. it receives the steering and speed values from the congatec module, which are transmitted to the freescale kl25z microcontroller, which in turn controls the rc-car model via pwm signals. 2. approach the development path of the approach to map the surroundings for an autonomous driving vehicle should be well-defined to avoid inconsistencies. the first step is answering the question which hardware will be used to detect objects. with the goal to use only a kinect camera as sensor, the approach relies on the processing of the depth, and standard images. both data represent a 2d respectively 3d model of the environment. the approach is characterised in three development steps: • 3d capturing, • projection, • object recognition. 28 http://dx.doi.org/10.14311/app.2016.6.0028 http://ojs.cvut.cz/ojs/index.php/app vol. 6/2016 mapping the surroundings as a requirement for autonomous driving 3d capturing obstacle detec on projec on and ltering kinect camera figure 1. spatial mapping of the current surroundings. it is developed to simplify complex processes, describe them in a proper way and visualize all of the main development steps in a manner that is easy to understand. according to the development steps, the approach is divided into three processes, see figure 1. at first, the kinect camera films the space in front of the vehicle in order to gather potential obstacles. the second step contains the projection of the depth images on a 2d ground plane. at last, the obstacles or object will be recognized during the third step. the three main steps will be explained below. 3. related work in 2002 sebastian thrun (see [5]) showed that the problem of terrain mapping is to generate a spatial model of a robot’s or vehicle’s environment using sensors and cameras. concerning this point there are several more subproblems: • measurement noise, • robot must choose its way during mapping process, • environment changes over time, • possibly the hardest problem — the so called data association problem: correspond sensor measurements to the same physical object in the real world. “mapping is largely considered the most difficult perceptual problem in robotics.” [5, p. 6]. in 1989 herbert et al. already presented the so called locus algorithm to build terrain maps based on a single active range sensor, by “computing the intersection of the surface observed by the sensor with the vertical line passing through (x,y)” [6, p. 998]. frankhauser et al. [4, p. 2] introduced an elevation mapping approach from a “robot-centric perspective” with relative distances to the robot and a permanent update process of the entire elevation map with information about the motion of the robot. kleiner and dornhege presented their autonomous driving vehicle for urban search and rescue (usar) with two different methods. first the rapid mapping of a large-scale environment by wheeled robots, and second the mapping of rough terrain by tracked robots [7]. hornung et al. [6] developed an open source library to represent the environment in memory using octrees. the cells of the discretized space can have one of the following states: unknown, free or occupied. to map the sensor data into the octree they use a probabilistic sensor model. thrun et al. [8] describes stanley, the winner of the 2005 darpa grand challenge. the robot was built by stanford university researchers and uses laser figure 2. overview of the kinect camera. range finders as well as radar sensors for sensing the environment. the evaluation of the sensor data is supported by artificial intelligence and machine learning algorithms. 4. 3d capturing as a 3d sensor is a rather expensive component, the kinect is a cheap alternative compared to the functions it provides. the camera is doing most of the calculations on its own soc, thus reducing the processing effort for the main unit. the first version of microsoft kinect camera was developed by prime sense, an israeli 3d sensing company, which was bought by apple in 2013. after that, microsoft created its own system based the existing prime sense technology [9]. 4.1. overview the kinect camera has an array of sensors, which can be divided into three categories: a depth sensor (green), a colour sensor (red) and an audio sensor (blue), as seen in figure 2. the audio sensor is not used for the 3d reconstruction. the colour sensor consists of a cmos camera with an infrared light filter. the main use of the colour sensor is the track detection. the colour image can be mapped with the depth information, although the resolution of the colour sensor is higher than that of the depth sensor. that is the reason why there are multiple colour pixels per one depth pixel. if there are more depth pixel, than colour pixels, the colour pixels need to be split into multiple smaller pixel, with the same information. the depth sensor consists of a cmos camera and an infrared led blaster. the cmos camera is filtered so that it just sees near-infrared light which is emitted by the infrared led blaster. all the information is processed in a specially designed soc (figure 3). this allows the soc to process a depth map containing all the depths for every pixel and map the colour pixel to it. the result of the processing is a data stream containing an audio, a video, and a depth stream. 29 m steininger, c. stephan, c. böhm et al. acta polytechnica ctu proceedings figure 3. 3d reconstruction of the kinect [11, 12]. 4.2. 3d reconstruction there are multiple ways of reconstructing the depth information from an image. the classic way needs two calibrated cameras at a defined distance. this allows for calculating the distance to the object with trigonometry formulas. the difficulty will be in differencing the objects resulting in the so called correspondence problem which “refers to the problem of ascertaining which parts of one image correspond to which parts of another image, where differences are due to movement of the camera, the elapse of time, and/or movement of objects in the photos.” [10]. the kinect overcomes this problem by using a method called “time of flight”. in this method the travel time of infrared light, which is emitted by a led, is measured for each pixel, as seen in 3. “a strobed infrared light illuminates the scene, the light is reflected by obstacles, and the time of flight for each pixel is registered by the infrared camera. internally, wave modulation and phase detection is used to estimate the distance to obstacles (indirect tof)” [13]. the problem is that the infrared light can be washed out by other near-infrared light sources like sunlight, resulting in the kinect not being able to detect the infrared light and losing the depth information. 5. projection and filtering the kinect camera is mounted on the chassis in an arbitrary fixed angle, tilted to the cars chassis. the exact angle in which the camera is mounted to the ground is not known. it can be measured, but to be independent from that setting, the ground plane coefficients, which are returned from the kinect, are used. the coordinate system of the returned 3d point cloud (as camera coordinate system in the following; indicated with a car subscript) does not match with the coordinate system of the car (as car coordinate system in the following; indicated with a camera subscript). in figure 4 both coordinate systems are shown. the body-frame which can be requested from the camera contains a four-dimensional vector representing the floor coefficients by ~pcoeff = (a; b; c; d). figure 4. the camera coordinate system (green) and the car coordinate system (blue) [11, 12]. these coefficients are combined in (1) to receive the plane equation for further mathematical operations: 0 = ax + by + cz + d (1) in a first step, the distance of every point to the ground is calculated. a point is considered to be a part of the ground if a certain limit is not exceeded. therefore they are removed from the point collection if the distance to the ground is smaller than a certain distance. the remaining points are processed in the following steps. in the next step, each frame of the kinect has a resolution of 512 by 424 and can thereby result in a large amount of 3d-pixels. a pixel is returned, if it is in the specified range of the camera. the nearest distance the camera returns a pixel for is about half a meter. the maximum distance is about eight meters. so there are quite a large number of points waiting to be computed. the rather small computing system with limited processing power on the car must handle this in real-time. for the presented project, a real-time reaction is defined with a latency of a few hundred milliseconds. to reach this goal, one issue is making the algorithm faster. this could be realised by: • removing as many points as possible to save energy and processing effort converting the points from one coordinate system to another, • adjusting processes how the data is processed step by step, • using a programming language with small processing overhead. to reduce the further processing effort the points get filtered next. a new parameter named mindistance is introduced in the gui (graphical user interface) where the points must have a greater distance from 30 vol. 6/2016 mapping the surroundings as a requirement for autonomous driving the camera than that value, to prevent recognising parts of the car as obstacles. this is needed due to the current setting of the car. there may be points above the height of the car so that the car can drive beneath. these points get removed if they are above a maximum height from the ground plane which is defined as a parameter named maxheight. obstacles far away are also not interesting for the process at the moment. in the same way points far away are also not interesting and get filtered with an allowed maximum distance named maxdistance. in this section some functions are introduced to simplify the explanation of the coordinate transformation. the first three equations deal with projections. project a vector v onto another vector vonto (2), also a vector v onto a plane e : 0 = (~x−~r) ·~n (3) and a point with position vector ~x onto a plane (4): fv2v(~vorig,~vonto) : ~vproj = ( ~vorig � ~vonto |~vonto| ) (2) fv2p(~vorig) : ~vproj = ~vorig −fv2v (~vorig,~n) ·~n (3) fp2p(~xorig) : ~xproj = ~xorig −fv2v (~xorig −~r,~n) (4) next the normalisation is defined in (5): fn(~vorig) : ~vnorm = ~vorig |~vorig| (5) since the source coordinate system of the kinect is in three dimensional space and the destination system is only two dimensional, it is necessary to project every detected 3d-pixel onto the ground plane. in the three dimensional coordinate space of the kinect we define a two dimensional coordinate space for the new coordinate system. the new coordinate system can be seen as a plane in the original space. therefore an origin and also two vectors for the two axes of that coordinate system are needed. the two dimensional coordinate system for the obstacle detection is implemented within the three dimensional coordinate system of the kinect camera. the camera position perpendicular above the ground is origin: ccamera = fp2p(o) with o as the origin of coordinate system. the axes are defined as ex = fn (fv2p ((0; 0;−1))) and ey = fn (fv2p ((−1; 0; 0))). in figure 4 you see the orientation of the axis. now, (6) can be used to project all points into the new car coordinate system: pcar = ( fv2v (pcamera −ccamera,ex) fv2v (pcamera −ccamera,ey ) ) (6) next, the obstacle detection will use these projected points and create a barrier along the nearest edge of the recognized obstacles. 6. obstacle detection when implementing algorithms for autonomous vehicles it is mandatory to detect and tag obstacles in front of the craft. as described, a depth image is 0 0.5 1 1.5 2 2.5 3 3.5 4 0 1 2 3 4 5 6 7 8 l o g a ri th m ic s c a li n g o f d is ta n c e distance (m) figure 5. logarithmic distance scaling for image size reduction purposes. provided by the kinect camera from where each single depth pixel is projected to a ground plane with each having the correct distance and position to the camera’s point of view during the first stage of image processing. as a first step, the depth points contributed by the kinect camera and the projection to the ground plane are transformed into a 2d image. the next three steps described in this section include: • distance scaling, • gaussian functions, • and the actual object and obstacle detection. 6.1. distance scaling the project team faced problems concerning computing power: the large number of pixels coming from a depth frame provided by the kinect camera exceeded the available and limited hardware. for the purpose of reducing the computing effort for processing the gaussian function (see subsection 6.2) the project team decided to scale the projected image’s size by its height logarithmically. as shown in figure 5, the scaling of distances allows to have more detailed image information in a closer range to the camera’s origin than in a larger distance to this specific point. the distance we get from the kinect camera is a floating point number, the index we need to reference a pixel within and on the image is an integer. if the distance is only rounded to the next integer value there is one pixel per meter. because of this a parameter is introduced which represents the resolution of the image’s distances. the scaled value is divided by this value to get the index as an integer, after rounding of this new integer we get the correct index value for the image. for example, 0.5 m in reality are mapped to an index value of 0.585 after the scaling, while 8 m in the real world, which is the kinect camera’s maximum recognition distance, correspond to an index value 3.171 logarithmically scaled. 31 m steininger, c. stephan, c. böhm et al. acta polytechnica ctu proceedings g(x,y) x y g(x,y) figure 6. gaussian distribution in 2d. as can be seen in figure 7 the opening angle of the camera is segmented with a specific resolution, which can be set as a parameter. every segment is mapped to a single rectangular column into the projected image. the advantage of this approach is that near distance pixels are richer in detail than more distant pixels. 6.2. gaussian function the next step is to smoothen the depth image to reduce single error pixels or noise and smoothen pixel clouds for further and more precise computations and calculations. although there are several options for smoothing images, the chosen solution is the gaussian filter g(x,y) = 1 2πσ2 e − x 2 +y2 2σ2 (7) in (7), the gaussian function for two dimensions is shown. where x is the distance from the origin on the horizontal axis and y is the distance from the origin on the vertical axis, and σ is the standard deviation of the gaussian distribution. the distribution can be seen in figure 6 for two-dimensional gaussian functions. this formula produces a surface of concentric circles with a gaussian distribution from the center point. the new value of each pixel is set to a weighted average of that specific pixel’s neighborhood. the original pixel’s value receives the highest gaussian value and neighborhood pixels receive smaller weights as their distance to the original pixel increases. this results in a blur effect [14, 15]. figure 7 shows three identical depth frames which are already projected to a ground plane (you see the frames in a top view) and transformed into a gray image. on the left-hand side, the single frame is displayed without any filtering or blurring. there are many false pixels, noise and distortion. a gaussian filter is already used in the middle frame, nevertheless quite a lot noise and other false pixels can be seen. the approach works in practice but still needs a little fine tuning. therefore a smoother configuration of the gaussian settings was used. the third frame on the right-hand side displays very little distortion, noise and false pixels, but provides clearly identifiable pixel clouds. these pixel clouds mostly represent objects or other obstacles in the range of the kinect camera. 1 2 3 figure 7. three identical frames with no gaussian filtering in frame 1, minimal gaussian filtering in frame 2 and for the project’s obstacle detection best gaussian filtering settings in frame 3. 6.3. detect obstacles the last step before actually being able to detect obstacles is removing unnecessary pixels which still are in the image after applying a gaussian filter. in the scope of the project a parameter was added to set a tolerance for still existing false or error pixels: threshold. it’s inevitable to do so, due to the fact that there are not only black and white pixels anymore (values 0 and 255) because of the gaussian filtering but also values in between. the final step is to find objects in the distancescaled and gaussian filtered image. after preparation of the given depth image with gaussian filtering and logarithmic distance scaling, the final step for finding and detecting objects in range is to find pixel clouds with a defined minimal distance (parameter in algorithm: mindistancebetweenobstacles in meters) to other clouds and therefore to other obstacles. when a pixel cloud (respectively an obstacle) is detected with no other obstacle in range of mindistancebetweenobstacles, it is marked as a detected obstacle. 7. conclusions the autonomous driving requires the continuous mapping of the surroundings. the presented approach and the hardware setting show that the mapping can be done with a kinect camera in the case of model cars in a scale up to 1 : 5. with the 3d capturing of the surroundings, the projection and filtering and obstacle detection, it is possible to get a temporary map of the car’s environment as a basis for the navigation process. the navigation itself relies on the a*-algorithm and uses the projected map including the detected 32 vol. 6/2016 mapping the surroundings as a requirement for autonomous driving obstacles as the digital weighted spatial information. in a next step, we are planning to test the whole configuration and the navigation process under different conditions, i.e. different surfaces or different weather conditions. the aim is to obtain feedback on the use and the performance of the stereo camera and the navigation algorithm. beyond the self-driving car, the presented approach can be implemented in other vehicles, like unmanned aerial vehicles (uav), autonomous underwater vehicles (auv) or even in robots for different services. references [1] a. m. kessler. elon musk says self-driving tesla cars will be in the u.s. by summer the new york times, 2015. [2] a. fisher. google’s road map to global domination the new york times, 2013. [3] a. geiger, p. lenz, r. urtasun. are we ready for autonomous driving? the kitti vision benchmark suite. in ieee conference on computer vision and pattern recognition (cvpr), pp. 3354–3361. 2012. doi:10.1109/cvpr.2012.6248074. [4] p. fankhauser, m. bloesch, c. gehring, et al. robot-centric elevation mapping with uncertainty estimates. in international conference on climbing and walking robots (clawar), epfl-conf-198746. 2014. doi:10.1142/9789814623353_0051. [5] s. thrun, et al. exploring artificial intelligence in the new millennium. vol. 1, chap. robotic mapping: a survey, pp. 1–35. 2002. [6] m. herbert, c. caillas, e. krotkov, et al. terrain mapping for a roving planetary explorer. in ieee international conference on robotics and automation, pp. 997–1002. 1989. doi:10.1109/robot.1989.100111. [7] a. kleiner, c. dornhege. real-time localization and elevation mapping within urban search and rescue scenarios. journal of field robotics 24(8-9):723–745, 2007. doi:10.1002/rob.20208. [8] s. thrun, m. montemerlo, h. dahlkamp, et al. stanley: the robot that won the darpa grand challenge. journal of field robotics 23(9):661–692, 2006. doi:10.1002/rob.20147. [9] primesense, 2016. page version id: 710008223. [10] correspondence problem, 2015. page version id: 661209488. [11] s. design. rc-xd (call of duty: black ops) 3d warehouse. [12] t. plus. kinect from microsoft 3d warehouse. [13] p. fankhauser, m. bloesch, d. rodriguez, et al. kinect v2 for mobile robot navigation: evaluation and modeling. in international conference on advanced robotics (icar), pp. 388–394. 2015. doi:10.1109/icar.2015.7251485. [14] m. s. nixon, a. s. aguado. feature extraction and image processing. newnes, oxford ; boston, 1st edn., 2002. oclc: ocm49045100. [15] l. g. shapiro, g. c. stockman. computer vision. prentice hall, upper saddle river, nj, 2001. 33 http://dx.doi.org/10.1109/cvpr.2012.6248074 http://dx.doi.org/10.1142/9789814623353_0051 http://dx.doi.org/10.1109/robot.1989.100111 http://dx.doi.org/10.1002/rob.20208 http://dx.doi.org/10.1002/rob.20147 http://dx.doi.org/10.1109/icar.2015.7251485 acta polytechnica ctu proceedings 6:28–33, 2016 1 introduction 2 approach 3 related work 4 3d capturing 4.1 overview 4.2 3d reconstruction 5 projection and filtering 6 obstacle detection 6.1 distance scaling 6.2 gaussian function 6.3 detect obstacles 7 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2018.20.0005 acta polytechnica ctu proceedings 20:5–9, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app btt & rflb the optimum set for steam turbine blades monitoring miroslav balda research and testing institute, tylova 46, 30100 plzeň, czech republic correspondence: balda@vzuplzen.cz abstract. btt, blade tip timing system, is a commercially available system generating files of precise times of blade tips when passing sensors attached in a machine stator. rflb, residual fatigue life of blades, is a postprocessor of those files evaluating estimates of fatigue lives of all blades fitted to the wheel. the set reduces a danger of unexpected blade failures. keywords: btt, rflb, fatigue, blades, maintenance. 1. introduction fatigue of materials is a very complicated phenomenon, which influences the quality of structures seriously. the bigger the structure is, the more important are the consequences of possible fatigue damages. the typical structures, in which any fatigue cracks should not occur, are turbomachines. any unexpected breakdown caused by the fatigue generates extreme losses of energy production. this is the reason why manufacturers of turbomachinery devote enormous care to a design of fatigue-save parts of a machine and to material testing. to common tests belong also fatigue tests of most exposed parts of the machine, blades of a turbine. the blades are working at rather complicated operational conditions like high radial stress from centrifugal forces and extreme velocity of wet steam turbulent flow exciting them to vibrations accompanied by alternating stresses. dynamic stresses at critical points of blades may initiate material damage, which results in a crack propagation when cumulated. since this phenomenon may end in a total breakdown of a blade followed by an unplanned shutdown of the machine, producers of turbomachines arm their products with systems, which are able to monitor blades at their critical places. 2. blade tip timing – btt the current monitoring systems used for that purpose are denoted as btt – blade tip timing systems because they are based on time measurement of blade tips when passing sensors (see fig. 1). a btt system consists of a set of sensors placed at the stator of the turbine above tips of blades of the selected wheel. outputs of ns sensors are connected to control inputs of their own counters of precise clock pulses. as soon as a tip of the vibrating blade passes a sensor, the sensor generates a pulse which initiates a transfer of the current content of the counter to its memory. there is still one sensor and the counter and its memory which counts pulses generated by a datum fixed at a shaft. at the end of the measurement, the contents of all memories are sent to ns + 1 files. the measured data, series of times, are processed offline into natural frequencies of blades and histograms of position magnitudes and other parameters of the measurement. 3. residual fatigue life of blades – rflb an application of a system rflb has in principle three main stages: • measurement of material properties, • modal analysis of the blade to be measured, • long time monitoring in situ in a power plant. all stages are very important for the final statement about the residual fatigue life of all measured blades. 3.1. material properties there are two material tests necessary for rflb system, the tensile test, and the high cycle fatigue test. the result of the tensile test is shown in the form of the diagram made by a testing device in fig.2. it may be accepted as a limiting case of the fatigue test during which the specimen has been broken at the half of the loading cycle. the high cycle fatigue (hcf) test is a different one. usually, a set of 8 10 specimens are loaded by harmonic processes of different amplitudes σam up to breaks after nam loading cycles [1]. the result of the test is a regression line to couples (nam,σam), called sn curve [2], which becomes a straight line when drawn in a diagram with axes log nam, log σam (see fig. 3). 3.2. modal analysis good information on the blade as an object of monitoring is very important. the detailed numerical analysis of the blade fitted in the wheel is made by a fem program. results of the analysis should contain the solution of the eigenvalue problem, e.g. natural 5 http://dx.doi.org/10.14311/app.2018.20.0005 http://ojs.cvut.cz/ojs/index.php/app miroslav balda acta polytechnica ctu proceedings figure 1. btt system with ns probes (sensors). frequencies and corresponding natural modes for a set of nodal diameter numbers. moreover, it is also important to have frequency response functions of circumferential deflection of a blade tip to stress tensors at critical places of the blade. 3.3. monitoring by the rflb the software system rflb complements the system btt by estimating fatigue lives of blades. it exploits the information contained in the btt output files for the purpose totally. such a complicated system, as the rflb is, should be split into modules fulfilling special tasks. the interactive matlab program rflb is built as a flexible one as possible with respect to events occurring in power plants. the system contains about 30 scripts and functions. the control module – rflb organizes, in a cooperation with the operator, the function of the system rflb. at first, the operator informs the module where to find the parameters of the blade, material, and results from the modal analysis. the operator may also interrupt the normal automatic run of the rflb when needed. there are many reasons for doing so. a typical situation occurs when the operator wants to look into the protocol or wants to plot the result of some past measurement. then, the program breaks the normal sequence of commands, fulfills the requirement of the operator, and continues just in the place, where has been interrupted. this behavior ensures the control module of the program. the module rflb initiates the run of the rflb with all new blades. it also enables to change some blades without a loss of continuity in the cumulation of damages of other blades. the same holds after a break in processing. the module btt is a driver for a continuous run of the system. it distributes tasks to specialized functions. the first of them is function getdata which reads the files generated by systems btt. there are served files from two btt systems, starman and hood. the function getdata is the only one which should be rebuilt if another btt system was applied for measurements. the reason for it is that the output files of systems btt are not standardized. the main role of function getdata is creating a vector composed of key-phasor times of once per revolution pulses from the red probe in figure 1 and a hyper-matrix (3d array) of times from all sensors and blade tips of the wheel. one vector of the 3d array is seen from figure 4. since the measured data are in the standard form, so the rest of the rflb system does not need any change. at the end of the function getdata, all-time series are fourier transformed into aliased fourier spectra by fast nieslony’s function rainflow [3]. the resulting fourier spectra are aliased in frequency due to the deep under-sampling of blade tip signals. frequency aliasing has to be eliminated by a spectrum reconstruction. this is implemented by function reconstr. it is well known that under-sampled signals defy reconstruction by normal way. fortunately, there is additional information at disposal. that is a set of natural frequencies obtained during the numerical 6 vol. 20/2018 btt & rflb the optimum set for steam turbine blades monitoring figure 2. material t671 tensile test. 0.5 1 10 100 1000 10000 100000 1e+06 1e+07 400 500 600 700 800 900 1000 1100 1200 1300 1400 1500 1600 1700 number of cycles n a s tr e s s a m p li tu d e σ a material t671 σ m =0 − woehler σ m 1 = 450 mpa σ m 2 = 600 mpa r m r 02 σ’ a0 σ’ a450 σ’ a600 σ c figure 3. t671 sn curves, σm = 0, 450, and 600 mpa. figure 4. vector of measured and reconstructed samples of one blade tip deflections. 7 miroslav balda acta polytechnica ctu proceedings figure 5. aliased dft of all blades in the wheel. modal analysis. the natural frequencies recalculated to the aliased ones serve for identification of evaluated peaks in dft. the frequency bands around dft peaks coinciding with aliased natural frequencies are moved on places corresponding real natural frequencies of the reconstructed spectra. in figure 5, there are positions of aliased natural frequencies shown in upper parts of the diagram by red ticks and order numbers. the upper frequency of the reconstructed fourier spectra is chosen by an order higher than the speed frequency, which is the sampling frequency fs, say, fs = 25 fs. the rest of the aliased spectra is spread within the reconstructed spectra under user wishes. the last stage of the reconstruction is the inverse fourier transform of the reconstructed spectra, what returns reconstructed signals in much higher resolution as seen in figure 4. if the moving of the spectrum bands was done with care, the reconstruction is successful, what is seen from figure 4. as soon as the reconstructed signals are ready, relative damages in critical places of the blades can be estimated by function damage. unfortunately, it is the most time-consuming operation, because of each blade signal has to undergo rain-flow analysis, what is a decomposition of each signal into full cycles, followed by a recalculation of time data into circumferential deviations and then to stress state in critical places on blades. for the purpose, the elements of the stress tensors obtained during modal analysis, are used for evaluating an effective (damaging) stress σd of kth loading cycle in the following form σ̃d,k,n = ∆tk × [ σ̃2x + σ̃ 2 y + σ̃ 2 z (1) − 2µ (σ̃xσ̃y + σ̃yσ̃z + σ̃zσ̃x) + k2c (τ̃ 2 xy + τ̃ 2 yz + τ̃ 2 zx) ]1/2 n , figure 6. relative damages of all blades. where kc =σc/τc ≈ √ 2(1 + µ) and symbols with tilde are dynamic components of a stress tensor in the critical point corresponding to nth mode of vibration. time-series of damaging stress course are decomposed into a set of full cycles and half cycles by a rainflow procedure getting thus stress amplitudes σak and mean stresses σmk needed for an estimation of elementary damage dk in kth cycle by a fatigue hypothesis. the most common one is the hypothesis of pålmgrenminer [4, 5] respecting an influence of σm [6]: dk,pm = 1/nam = 2 ( sa s′a0 − sm )w , (2) where fictive amplitude s′am replaces either σ′am or 8 vol. 20/2018 btt & rflb the optimum set for steam turbine blades monitoring τ′am. it is evident from figure 3 that its values are given by the positions of intersections of a particular sn curve with the line nam = 0.5. for all family of sn curves holds that s′am = s′a0 − sm. this property of sn curves of steels saves time and money because it canceled the necessity of making expensive tests of prestressed materials. the system rflb cumulates elementary relative damages defined by equation (2), so it is able to yield information on the total relative damage d ≤ 1 of any blade of the wheel to the operator at any moment. in case, the damage d reached a chosen amount, the system issues an alarm message. residual fatigue life, defined as l = 1 − d, says how much time still remains to break down. 4. conclusions the article has shown the advantages gained by a combination of a hardware system btt with the software system rflb. installation of the set of both items in power plants for monitoring blades in potentially uncertain turbine stages may be very effective for the reliability of a machine because it diminishes a danger of unexpected breakdowns and enables to plan times of maintenance and a number of spare blades in store. list of symbols dam relative damage of a cycle with σa and σm d total relative damage fn natural frequency [hz] fs sampling frequency [hz] fs sampling frequency of reconstructed spectra [hz] fn nyquist frequency = fs/2 [hz] l relative residual fatigue life nam number of cycles to failure under σa and σm rm material strength [mpa] rp yield limit [mpa] s either σ or τ [mpa] t time [s] ts sampling period = 1/fs [s] w exponent of woehler’s curve µ poisson’s ratio σ stress in tension [mpa] σa amplitude of a stress cycle [mpa] σm mean stress of a stress cycle [mpa] σam amplitude of a cycle by mean stress [mpa] σ′am intersection point of sn curve with n = 0.5 [mpa] τ stress in torsion [mpa] acknowledgements the author appreciates a kind support of his research from the technological agency of the czech republic via the center of competence no. te01020068, center of research and experimental development of reliable energy production references [1] j. chvojan: a test of material t671 high cycle fatigue (in czech), vzú plzeň, testing record no. 846/2003 [2] čsn 42 0368: fatigue testing of metals. statistical evaluation of fatigue test results of metals (in czech), czech normalizing institute, prague, 1973. [3] a. nieslony: rainflow counting algorithm. mathworks, file exchange, file no. 3026, 2003 [4] a. pålmgren: die lebensdauer von kugellagern, vdi zeitschrift (68), pp. 339-341, 1924. [5] m.a. miner: cumulative damage in fatigue, trans. asme, journal of applied mechanics, (12) pp. a159-a164, 1945. [6] m. balda: fatigue life assessment of prestressed material, engineering mechanics, c. fischer & j. náprstek eds., itam cas cz, proc. conf. engineering mechanics, svratka, may 14-17, 2018 9 acta polytechnica ctu proceedings 20:5–9, 2019 1 introduction 2 blade tip timing – btt 3 residual fatigue life of blades – rflb 3.1 material properties 3.2 modal analysis 3.3 monitoring by the rflb 4 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.21.0033 acta polytechnica ctu proceedings 21:33–37, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app reconstruction of the monastery next to the prague castle jakub řeháka, ∗, eva burgetováa, josef řehákb, stanislav řehákb a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b speleo-řehák, úvoz 6, 118 00 prague 1, czech republic ∗ corresponding author: jakub.rehak.1@fsv.cvut.cz abstract. the paper deals with a rehabilitation of the historic building of carmel monastery next to the prague castle. a structural survey of building was carried out and the historical drainage and ventilation system was identified. based on structural survey in-situ the reasons of failures were analyzed and remedial works of historic monastery were proposed. keywords: static failures, rising damp, structural survey, historic drainage systems, natural ventilation, sewer, rehabilitation. 1. introduction protection of buildings from water and moisture relates to the most important measures in ensuring the service life and serviceability of buildings. of great importance, as well, is structural solution of the designed rehabilitation procedure. unqualified interventions and measures often lead to failures and in the final phase, the deterioration of the whole complex. for this reason, the problems of long-term efficiency and reliability of rehabilitation methods are becoming a subject of great interest, both in terms of research and practical applications. the focus is mainly on listed buildings where technological designs are the most complicated. figure 1. carmel monastery in the prague castle square, view to the southern terrace of monastery. 2. monastic water systems complex water systems in the high middle ages was a phenomenon that occurred throughout europe. construction of new complex hydraulic systems seems to have begun in the 11th century, and it became increasingly widespread during following centuries. the concentration of water systems looks higher in germany than elsewhere. the rapid and widespread diffusion of the orders was followed by a rapid and widespread diffusion of hydraulic technology [1]. this technical knowledge of water and drainage systems was brought to bohemia primarily by cistercians and benedictines as a part of the spread the creed in the 13th century. the monastery of discalced carmelite nuns (the carmel of saint joseph) is located on the prague castle square. a part of the monastery grounds is the st benedict’s church, first mentioned in 1353, later the church underwent a number of reconstructions. in 1626 the monastery complex was added to it. its present form dates back to the middle of the 18th century. the carmel monastery was built on a steep south side of the castle promontory (fig. 1 and 4). the subsoil consists of weathered clay schist, which are unstable and slushy. in the middle ages a number of drainage and ventilation systems had served to prevent penetration of moisture into the structures, they drained water and dampness out of the structures and maintained them reasonably dry [2]. the baroque monastery buildings were forcibly abandoned in 1950 and the entire complex was costly adapted for use as a luxury hotel. between 1985 and 1991, the monastery underwent extensive reconstruction and was returned to its original religious purpose in 1992. figure 2. carmel monastery view of prague by f. van ouden allen, 1685. however, serious static problems manifesting by cracks have appeared and the building was moving slightly downhill to the street radniční schody (townhall stairs). according to the report based on visual 33 http://dx.doi.org/10.14311/app.2019.21.0033 http://ojs.cvut.cz/ojs/index.php/app j. řehák, e. burgetová, j. řehák, s. řehák acta polytechnica ctu proceedings inspection it was proposed to stabilize the building by grouting of the foundations and by fixing with reinforced ring beam and ties (fig. 3). the costs of remedial measures were estimated at almost czk 100 million, which was above the possibility of the nuns, as the owner of the building. figure 3. carmel monastery: proposed remedial measures. 3. experimental methods therefore carmelite nuns have addressed speleořehák company, specialized in rehabilitation of drainage systems and associated failures in historic buildings. during the subsequent structural surveys serious problems, that were causing the stability disruption, were revealed. figure 4. location of the monastery on a steep south side of the castle promontory between prague castle square and radniční schody (town hall stairs). historic drainage systems, however by ignorance of their importance and function, were clogged or filled up during reconstruction, possibly walled up as unnecessarily due to amateurish or inadequately carried out surveys or due to ignorance or negligence of geological and hydrogeological conditions. lessons from the functioning of these systems and their importance are unfortunately often acquired negatively during the laborious repairing of damage caused by undue interference in their functions. stability disruption of masonry of the carmel monastery causing its degradation was the result of interaction between multiple negative effects listed below: • leaking pipes the first defect was diagnosed in the main courtyard of the monastery. after realization of the borehole into the damaged masonry and the following extraction of the wet fill, the cause of the problem was discovered a leaking waste pipe. the water from the pipeline flowed through the backfill and saturated the adjacent structures with water. • closure of the main gallery at the same time, by the drill hole the rest of a lateral gallery was discovered. the main gallery, according to later found historical documentation, led originally from fountain in the courtyard, but later was closed and abolished. the whole area, as it turned out, belonged to the historic drainage system of the monastery and was covered with earth and debris. from this part of underground, about 70 tons of building waste probably from the time of reconstruction in the 1990s before the return of the monastery to the nuns, was removed. • collapse of the gallery walls the side walls of gallery have collapsed due to lack of maintenance of the underground system. this resulted in the vault falling in and the cavern originated here. if this situation is not solved, it could threaten imminently the surrounding buildings. • break of the toilet pit in the main gallery, fully filled up with rubbish, under monastery corridor the toilet (latrine) pit was discovered. the pit originally led up to the first floor, where it gradually moved into the size of a larger chimney and was led up to the roof. thanks to the chimney effect, this shaft provided sufficient air dynamics to help drying of the entire building through the drainage system. the chimney effect produced a negative pressure in the system, ensuring the odor from the pit will remove above the roof and at the same time preventing the odor from entering the building. during the reconstruction of the monastery, after the expulsion of the nuns in the 1950s, this toilet pit was interrupted on the ground floor by the corridor, thus preventing natural air flow and effective drying of the building. figure 5. carmel monastery: cross section of the main gallery with marks of the caverns. • exposed footing bottom 34 vol. 21/2019 reconstruction of the monastery next to the prague castle behind the toilet pit below the perimeter wall a large cavern with a diameter of 2.5 3 meters was discovered. looking from the gallery to the cavern, the footing bottom of perimeter wall in length of about 2m is suspended ”in the air” (without contact with subsoil). exposed footing bottom could not fulfill its function of transferring the load to the subsoil. the reason for the cavern was the total neglect and subsequent destruction of the underground system in combination with insufficient rainfall down pipe during the aforementioned rebuilding. water leaking from rainfall pipe saturated soil resulted in form of a cavern under the footing bottom. • the gallery clogged with concrete in 2013, due to removing of waste fill and rubbish, new part of historic underground drainage system was found. newly discovered gallery led up to the edge of the retaining wall of the monastery terrace. however, the gallery was filled up with concrete at the level of the retaining wall foundation. it is clear, that the drainage system continues further down below the radniční stairs and ensures the drainage of moisture and water also for the houses laid down below, where the survey subsequently confirmed this assumption. by blocking the gallery, the system could not perform its function. in addition, before the concrete ”plug”, leaking sewage pipe is connected to the gallery, which explains the critical state of the plaster on the supporting wall of radniční stairs. spalling of the plaster was caused by a leaking pipeline that supplied backfill by water, so it functioned as a water reservoir and constantly subsidize the support wall with moisture. the survey carried out in the carmel monastery showed, that moisture imminently threatened construction and represented a heavy burden for the entire building. 4. the proposal for redevelopment there are in principle two possible approaches for proposal of moisture redevelopment: • to fully respect and rehabilitate the original historical moisture proofing methods based on the ventilation of moist air from the structures, while assessing their options in relation to the existing use of the building. the results of applying these methods will take effect in the long term and require careful maintenance [3]. • to use of the latest technologies and materials, which are a perfect for sealing surfaces (e.g. grouting, waterproofing sealants etc.). this method is fast, the results will come immediately after application, but this is an irreversible action on the historic building [3]. restoration of traditional systems of protection of buildings from water and moisture should be prifigure 6. carmel monastery: radniční schody (town hall stairs) spalling of the plaster due to water-saturated backfill, the gallery clogged with concrete. oritized over unsuitable contemporary methods or costly technological measures. based on the results of the survey and analysis of the failures of the carmel monastery the rehabilitation procedures that pose minimal intervention into the structure, were designed in the context of the entire historic area. with an emphasis on the heritage value of the building it was used the original historical technologies for removal of moisture based on the principle of natural venting dampness from the structures. water-saturated backfill was manually removed, sewer pipes were repaired, opened up and unwanted ground water discharged from the building. the connection to a ventilation shaft was completed. moisture from the gallery is now diverted into the cleaned ventilation shaft and due to the chimney effect permeates up through the building above the roof. at present, the galleries of the old system have been mended and they serve as drains and part of the measures securing the building from moisture. 5. evaluation of remediation currently, it can be stated that the drainage and ventilation system is functional and moisture regime in building is stabilized. realized remedial variant has proved to be functional and effective within three years it has completely succeeded in drying the walls. the process of plastic deformation with cracking was stopped, cracks were bonded and sliding stabilized. it is clear, that these measures are sufficiently effective to prevent any further development of former structural problems, no other structural intervention were needed. 35 j. řehák, e. burgetová, j. řehák, s. řehák acta polytechnica ctu proceedings figure 7. carmel monastery: reconstructed main gallery. investigation involving the testing of rehabilitation methods and procedures in existing buildings is apparently the most reliable approach for obtaining conclusive results on the effectiveness of damp-removal measures. the experimental testing is useful not only for the assessment of the efficiency of remedial technologies, but also for a deeper investigation of moisture transport phenomena [4]. this appears an attractive way for better approaching the real features of building materials and, at the same time actual climatic conditions experienced on site. the phenomenon of moisture movement is extremely variable from case to case, depending on the specific features of the masonry: direction, quantity and continuity of water supply from ground, depending on seasonal and accidental factors, evaporation condition, thickness and arrangement of the bricks in the masonry, nature of mortar joints etc. [4]. in 2019 the long term monitoring supported by the european structural and investment funds, op prague: analysis of feasibility, concept dk-07, ”measurement system for tests and verification of the algorithm for description of properties and functions of the historic conduits” will verify the remedial measures carried out in the carmel monastery. testing of the moisture transport in historic drainage and ventilation system in the carmel monastery will use climatic and microclimatic data which allowed estimating quite accurately the indoor and outdoor evaporation. the efficiency of the air system will be monitored through the velocity and direction of airflow in the system. the collected data could allow the visual presentation of the behaviour of treated structures, its prediction and could help in the design of remedial measures as well. 6. conclusions the survey of the monastery gave an overview of the hydrogeological situation of this locality and identified the historic drainage and ventilation system. the solution of historic waterproofing was found and respected and based on its effectiveness was rehabilitated. crucial places, where water accumulated, were eliminated by opening up the gallery and cleaning the shaft. the sewer pipes were repaired completely. thus at the same time air circulation was restored and the high air humidity removed from the building. underground drainage systems are important and inseparable part of the historic monument, to complement its building history and have for its existence and preservation of its irreplaceable function. in historical buildings, it is desirable within the investigation of moisture conditions to identify, properly understand and restore the frequently faulty or damaged original insulation, ventilation or drainage systems, which was based on many years of experience of traditional building [5]. the understanding of the original or older solution, its technical parameters and performance is the first step towards finding out the causes of potential failures of the building and their removal allowing, at the same time, restoration of the function of the waterproofing system [6]. the reconstruction of historical drainage systems belongs to the basic tasks of conservation of buildings. experience of numerous cloister complexes show that the principal cause of increased ground moisture content and changes in subsoil manifested through structural failures of buildings is the breakdown of older drainage and ventilation systems [5]. the case of the carmel monastery confirmed that certain static failures and defects can be solved by additional treatment of moisture regime in building subsoil only and the application of invasive static remedial interventions is not necessary. at present, the galleries of the historic system have been mended and served as drains and sliding accompanied by cracking was prevented. the long term measurements of moisture movement verifying the remedial measures are planned in 2019. references [1] r. magnusson. water technology in the middle ages: cities, monasteries, and waterworks after the roman empire. baltimore: johns hopkins university press, 2001. [2] j. řehák. speleological survey of schwarzenberg palace praha 1, hradčanské nám. čp. 185. 2000. speleo-josef řehák, archiv speleo. [3] e. burgetová, k. michalová. a survey of historic galleries in the area of the prague castle. in advanced materials research, vol. 1122, pp. 61–64. trans tech publ, 2015. [4] e. franzoni. rising damp removal from historical masonries: a still open challenge. construction and building materials 54:123–136, 2014. 36 vol. 21/2019 reconstruction of the monastery next to the prague castle [5] e. burgetová. analysis of rehabilitation methods for rising damp in masonry 2005. habilitation thesis. [6] j. řehák, e. burgetová, j. řehák. moisture survey of former convent in the vicinity of prague castle 2017. engineering material. 37 acta polytechnica ctu proceedings 21:33–37, 2019 1 introduction 2 monastic water systems 3 experimental methods 4 the proposal for redevelopment 5 evaluation of remediation 6 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0067 acta polytechnica ctu proceedings 22:67–71, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app verification of the basic properties of recycled and natural aggregates karel mikulica∗, iveta hájková brno university of technology, faculty of civil engineering, veveří 331/95, 602 00 brno, czech republic ∗ corresponding author: mikulica.k@fce.vutbr.cz abstract. in the future, it is planned to use up to 50 % of construction and demolition waste (c&dw) for the production of new building structures. this leads us to think about how we can use recycled concrete aggregate (rca) as a substitute for natural aggregate (na) in concrete mixtures. this is why we compare the two typical representatives of recycled aggregates with a representative of natural aggregates. as a representative of recycled aggregates, we chose pure concrete recycled from the cutting of concrete and mixed recyclate from the demolition of the apartment building. as a representative of natural stone, we chose the extracted aggregate. keywords: construction and demolition waste (c&dw), recycled concrete aggregate (rca), natural aggregate (na), stone testing. 1. introduction aggregates form the supporting framework of the concrete mixture, occupying almost 80 % of the volume in the concrete. because of this, the issues of sustainability of construction and recycling come to the subconscious. although there is currently a very rapid development in the field of building materials, concrete remains a building that has one of the widest applications in the construction industry. the composition of the concrete can be very varied depending on the filler, additives or additives used. however, the filler function always forms aggregates. depending on the properties of the structures, it is possible to modify the concrete composition to meet individual requirements, eg strength, compactness, resistance to the influences. one of the requirements may also be its subsequent recycling. concrete as a composite material consists of two components, aggregate and cement slurry. however, the third component, which is a layer at the interface of the cement slurry and aggregate, is thin, about 10 50 microns, and we call it the transit zone. it is formed by the fact that the grains of cement are distant from the grain surface of the aggregate and thus the grain is covered only with a layer of water and a small amount of cement particles. the effort of cement to deviate from the aggregate is also supported by the homogenization process in the mixer. [1] plain concrete from recycled concrete does not reach the original concrete properties. recycled concrete often has less strength and more creep and shrinkage than the original concrete. improvement of these properties can occur with the use of epoxy dispersions without the presence of a solvent. this improves the adhesive properties of cement bonding cement recyclates and, if the dispersions are applied to dry recycled material, have an effect on reducing the water absorption value. the characteristics of concrete from recycled aggregate are affected by impurities that contain recycled aggregate. using the dirt will be transferred to a new material or construction. these are impurities in the form of clay, glass, plaster, plastics, textiles, papers, etc. the importance of dirt and their percentage of recycled concrete is not negligible, reflecting the quality of the newly developed material. [2] the use of modified construction and demolition waste (c&dw) as a partial substitute for natural aggregates (na) saves not only natural resources, but at the same time contributes to the reduction of the amount of construction waste in landfills. reasons for using recycled aggregate (rca), however, are not only these stated benefits that correspond to the objectives of sustainable construction, but also its cost and affordability, which is more favorable in many states than natural aggregates. the disadvantage of recycling, however, is that secondary material no longer achieves the precise properties of the primary raw material, in our case, natural aggregate. methods of testing, sorting and classification are covered by european standards en 12620 + a1 aggregates for concrete [3] and en 933-11 testing of the geometrical properties of aggregate classification of components of gross recycled aggregate. [4] therefore, it is not used to a greater extent. an important aspect in the production of recycled aggregate is also the economic aspect, it is important that the production of recycled aggregate is not more expensive than the extraction of natural raw materials, thus the importance of recycling would be lost. because the durability of the concrete is limited, we need to anticipate its decay and subsequent demolition as soon as a new concrete structure is built. therefore, we should think about the solution of the concrete afterwards when it stops performing its function. 67 http://dx.doi.org/10.14311/app.2019.22.0067 http://ojs.cvut.cz/ojs/index.php/app karel mikulica, iveta hájková acta polytechnica ctu proceedings it is, however, appropriate to say that rca has not yet been widely used in the construction industry due to the diversity of individual recycled concrete mixtures. for this reason, several physico-mechanical properties such as bulk density, porosity, water absorption, frost resistance, reactivity of unhydrated cement in rca, leachability of pollutants, purity of rca itself, alkaline silicate reaction, etc. are observed in rca. [5] 2. materials and methods to verify the basic properties of recycled aggregates, we chose two types of slightly different concrete recyclates. rca 1 is the concrete recycled from the cutting of the outdoor concrete surface. on the other hand, rca 2 comes from the demolition of a residential building and contains a wide range of different materials from the concrete ceramic shards, glass pieces. as a reference natural stone for our needs was selected directly from the water extracted rock. in order to achieve the same input conditions for all the aggregates compared, we chose the gravel fraction of 0 16 mm for testing. we subjected to the tests and measurements described below. grain. determination of granularity, or screening of aggregates, is one of the basic tests performed on aggregates according to en 933-1 testing for particle size distribution sieving method. [6] the principle is to sort and separate the material by means of a set of sieve into several fractions with a descending size of holes. the result is graphical representation of the representation of individual fractions in the aggregate sample. the sample size of the test specimen depends on the maximum grain size d. in our case, we used a fraction of 0 32 mm for all tested aggregates. at the same time, we determined the percentage of fine particles < 0.063 mm in the form of nozzles. density. the rca bulk density depends on the residual mortar. since the residual mortar has a lower density than aggregate, the bulk density of the recycled aggregate is lower than natural aggregate. its value is from 2100 2500 kg/m3, natural aggregate has a bulk density of about 2400 2900 kg/m3. other dependencies are grain size. the smaller the grain, the lower the bulk density, but also the strength of the concrete. bulk density has an effect on water absorption, resistance to chrl. the smaller the concrete is more absorbent and less resistance to chrl. [5] testing of the bulk density of aggregates as described in en 1097-6 tests for mechanical and physical properties of aggregates part 6: determination of particle density and water absorption. [7] suffocation. the surface of the recycled aggregate is dusty and cracked, causing bumps and crushing when recycling concrete slabs. the cement is more porous and more absorbent in the area of the transit zone, leading to the formation of larger ca(oh)2 particles and thus the increased permeability of the concrete. [5] therefore, we tested 24 hours of water absorption on the samples according to en 1097-6 tests for mechanical and physical properties of aggregates part 6: determination of particle density and water absorption. [7] shape index. this test is determined only for rough aggregate according to en 933-4 tests for geometrical properties of aggregates part 4: determination of particle shape shape index. [8] principle is the distribution of individual grains according to the ratio of their length to thickness (l/e), where the shape index is calculated as a grain fraction with a ratio greater than 3. bulk density. this is the mass of the aggregate unit of aggregate with cavities and pores, including spaces and between grains. we distinguish bulk density of loose and shattered aggregate. for this post, we tested the first property we determined according to en 1097-3 tests for mechanical and physical properties of aggregates part 3: determination of loose bulk density and voids. [9] 3. results this part of the paper summarizes the measured results of our tested samples of two recycled concrete aggregates (rca 1 and rca 2) with one reference mined aggregate (na reference). the results are written in the enclosed tables. in the first step, we first determined the percentage of fine particles, when this test went freely to the screening of the test samples. this basis specification is introduced in table 1. the detailed particle size distribution of studied materials in shown in table 2. resulted gradation curve is subsequently depicted in figure 1. a further test was the comparison of volumetric weights to avoid variations due to the different moisture content of the aggregates, we determined the bulk density on samples dried to a steady state. the results obtained are summarized in table 3. we used samples from the bulk density test to determine the 24-hour water absorption so that the conditions for all samples were the same. the results of water absorption test are shown in table 4. the shape index was determined on a limited fraction of 4 8 and 8 16 mm using a special sliding scale with a ratio of 1:3. the result is an l/e ratio where grains with a greater than 3 ratio are not suitable for use in the concrete mixture due to the unevenness of the material, which could occur during compaction of the concrete mixture to form caves or to break these non-conforming grains. the detailed results are introduced in table 5. the final test was to determine the bulk density of loose aggregate using a standard vessel of known volume, in our case 10 liters. obtained results of single studied aggregates are well visible in table 6. 68 vol. 22/2019 verification of the basic properties of recycled and natural aggregates sample identification percentage of fine particles (< 63 µm) [%] reference na 0.23 rca 1 0.45 rca 2 1.38 table 1. determination of percentage of fine particles. mesh sizes [mm] total fall [%] reference na rca 1 rca 2 0.063 0.0 0.0 0.0 0.125 0.1 0.2 1.6 0.25 0.6 0.7 5.2 0.5 2.9 1.4 15.1 1 6.2 2.0 25.9 2 9.4 2.3 34.4 4 14.3 2.5 44.2 5.6 23.3 3.1 51.1 8 50.9 14.7 58.4 11.2 76.8 58.2 65.3 16 98.1 97.7 71.7 22.4 100 100 76.1 31.5 100 100 85.6 table 2. the results of the screening analysis of aggregates. sample identification faction [mm] density of the dried sample ρrd [kg/m3] reference na 0 – 4 26304 16 2580 rca 1 0 – 4 20804 16 2270 rca 2 0 – 4 22404 16 2230 table 3. determination of bulk density in the dried state. figure 1. graphical determination of the results of screen analysis of aggregates. 69 karel mikulica, iveta hájková acta polytechnica ctu proceedings sample identification faction [mm] weariness after 24 hours wa24 [%] reference na 0 – 4 1.24 16 0.75 rca 1 0 – 4 9.54 16 5.7 rca 2 0 – 4 5.94 16 5.3 table 4. determination of 24 hours water absorption. sample identification faction [mm] shape index si reference na 4 8 108 16 13 rca 1 4 8 108 16 4 rca 2 4 8 68 16 9 table 5. determination of the shape index. sample identification faction [mm] the bulk density of the loose sample [kg/m3] reference na 0 16 1580 rca 1 0 16 1120 rca 2 0 16 1550 table 6. determination of bulk density of loose aggregate. 4. conclusion the largest proportion of fine particles and a significant fraction of particles below 4 mm had rca 2, the resulting value being about 3 times higher than for rca 1 and even 6 times higher than for na reference. the screen analysis itself showed that rca 2 contains almost 45% of grains below 4 mm, whereas rca 1 has only 2.5%. this result was probably due to the different jaw setting of the crushing machine. determination of bulk density has shown that both recycled concrete aggregates have a similar bulk density of approximately 2250 kg/m3, ie about 350 kg/m3 lower than the reference na 2600 kg/m3. 24 hours of water absorption has highlighted the most common problem in using recycled concrete aggregates, ie water absorption. this is about 5 times higher for recycled concrete and for rca 1 fraction 0 4 mm even 9 times higher than for reference na (about 1%). as a result, a higher proportion of cement dust may be present in the fraction. when evaluating the shape index test best suited to 8 16 mm rca 2 fractions with 4, we can see why the lowest bulk density of 1120 kg/m3 was measured for rca 1 compared to approximately 1550 kg/m3 for reference na and rca 2 . acknowledgements this work was financially supported by project fast-j18-5554 “possibilities of using recycled concrete for reuse in construction”. references [1] a. shah, b. r. chughtai, m. m. ghumman, et al. mechanical properties of recycled aggregates concrete, concrete engineering. https:// www.engineeringcivil.com/mechanical-propertiesof-recycled-aggregates-concrete.html. [2] t. pavlů. zkoušení a vlastnosti recyklovaného kameniva pro použití do betonu, stavba.tzb-info.cz 10265. https://stavba.tzb-info.cz/beton-malty-omitky/ 10265-zkouseni-a-vlastnosti-recyklovanehokameniva-pro-pouziti-do-betonu, 2013. [3] en 12620+a1 aggregates for concrete. european standard, european commitee for standartization. [4] en 933-11 testing of the geometrical properties of aggregate classification of components of gross recycled aggregate. european standard, european commitee for standartization. [5] e. skriňáková. studium možností využití kameniva z recyklovaných betonů pro výrobu konstrukčních betonů. bachelor thesis, vysoké učení technické v brně, fakulta stavební, ústav technologie stavebních hmot a dílců, brno, 2014. [6] en 933-1 tests for geometrical properties of aggregates part 1: determination of particle size 70 https://www.engineeringcivil.com/mechanical-properties-of-recycled-aggregates-concrete.html https://www.engineeringcivil.com/mechanical-properties-of-recycled-aggregates-concrete.html https://www.engineeringcivil.com/mechanical-properties-of-recycled-aggregates-concrete.html https://stavba.tzb-info.cz/beton-malty-omitky/10265-zkouseni-a-vlastnosti-recyklovaneho-kameniva-pro-pouziti-do-betonu https://stavba.tzb-info.cz/beton-malty-omitky/10265-zkouseni-a-vlastnosti-recyklovaneho-kameniva-pro-pouziti-do-betonu https://stavba.tzb-info.cz/beton-malty-omitky/10265-zkouseni-a-vlastnosti-recyklovaneho-kameniva-pro-pouziti-do-betonu vol. 22/2019 verification of the basic properties of recycled and natural aggregates distribution sieving method. european standard, european commitee for standartization. [7] en 1097-6 tests for mechanical and physical properties of aggregates part 6: determination of particle density and water absorption. european standard, european commitee for standartization. [8] en 933-4 tests for geometrical properties of aggregates part 4: determination of particle shape shape index. european standard, european commitee for standartization. [9] en 1097-3 tests for mechanical and physical properties of aggregates part 3: determination of loose bulk density and voids. european standard, european commitee for standartization. 71 acta polytechnica ctu proceedings 22:67–71, 2019 1 introduction 2 materials and methods 3 results 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0099 acta polytechnica ctu proceedings 15:99–103, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app characterization of cement-based composite exposed to high temperatures via ultrasonic pulse method iva rozsypalová∗, michal vyhlídal, richard dvořák, tomáš majda, libor topolář, luboš pazdera, hana šimonová, zbyněk keršner brno university of technology, faculty of civil engineering, veveří 331/95, 602 00 brno, czech republic ∗ corresponding author: iva.rozsypalova@vutbr.cz abstract. in this paper, the attention is paid to the investigation of the influence of high temperature acting on specimens made from specially designed cement-based composite. the experimental programme was carried out on six sets of beam specimens with the dimensions of 20 × 40 × 200 mm. the specimens were loaded to a pre-set temperature of 100, 200, 400, 600, 800 and 1000 ◦c and then the temperature was kept for 60 minutes. when the temperature loading had been done, the specimens were left to cool down to the ambient temperature. after that, the ultrasonic pulse method was used to determine the degree of damage of temperature loaded specimens. the measured data obtained by this non-destructive method are in high correlation with values of informative compressive strength of the composite obtained after the temperature loading of specimens. keywords: fine-grained cement-based composite, glass spherical aggregate, high temperature, ultrasonic pulse method. 1. introduction cement-based composites belong to the frequently used building materials for a wide range of applications [1, 2]. structures and construction components utilizing this composite were commonly designed assuming normal service temperatures. however, high temperatures acting on cement-based composite – for example during the fire etc. – cause a wide range of physical and chemical processes, which result in changes in the structure of composite and thus affects its mechanical properties [3, 4]. according to the theory of fracture mechanics, cement-based composites, in this case fine-grained cement-based composite, show quasi-brittle behaviour. specimens made from these materials have the ability to carry the load even after the deviation from a linear branch of load–displacement diagram until the peak point and then the decrease of the loading force follows until the failure – see e.g. [5, 6]. one of the reasons for this behaviour could be the existence of the interfacial transition zone (itz). this well-known zone around aggregate particles of a few micrometers in size has the specific features. these are mainly higher porosity and calcium hydroxide contents compared to the bulk matrix [7–12]. the scope of this paper is to show the results of a pilot study which deals with the characterization of the damage degree of special cement-based composite specimens exposed to high temperatures via ultrasonic pulse method. 2. experimental programme – materials, specimens, methods three-part silicone moulds were made in plywood frame and subsequently used to produce the test specimens, see figure 1. the test specimens were manufactured from the special designed fine-grained cementbased composite. the fresh mixture was made using the spherical sodium-potassium glass aggregates with a diameter of 2 ± 0.2 mm, portland cement cem i 42.5 r (from cement plant mokrá) and water in a ratio of 3:1:0.35. the components were mixed in laboratory conditions using the standard mixer. after pouring and compaction of the fresh mixture, the moulds was sealed with a thin pe foil and placed in the stabilized laboratory conditions for 1 day, see figure 2. after demoulding the test specimens had been stored in a water bath until their testing. the experimental programme was carried out on six sets of specimens according to considered maximum temperature. always, three test specimens with the nominal dimensions of 20 × 40 × 200 mm were prepared per each testing set. a laboratory furnace classic type 5013 with 3 sides of heating coils was used to heat of test specimens. the temperature loading started at the laboratory temperature of 20 ◦c. the specimens were loaded with the temperature rate of 5 ◦c/min up to a pre-set temperature. the pre-set temperatures were 100 ◦c (reference temperature), 200 ◦c, 400 ◦c, 600 ◦c, 800 ◦c, and 1000 ◦c. the next part of the temperature–time curve was constant, i.e. the maximum temperature was kept for 60 minutes. after that, the specimens were left to cool down to the ambient temperature and had been placed in a room 99 http://dx.doi.org/10.14311/app.2018.15.0099 http://ojs.cvut.cz/ojs/index.php/app i. rozsypalová, m. vyhlídal, r. dvořák et al. acta polytechnica ctu proceedings figure 1. illustration of silicone moulds preparing. figure 2. illustration of manufacture of test specimens. with the temperature of 20 ◦c. all specimens were tested by ultrasonic pulse method before and after the temperature loading performed at the laboratory temperature. during the ultrasonic pulse method measurement, the specimens were excited by a sensor placed in the centre of side cross-section of specimen – transmitter t (see figure 3). impulses were generated by the pulse generator agilent 33220a. the acoustic emission system xedo was used to detect the waves. three piezoelectric sensors were placed on different locations of the specimen surface, see figure 3. the first sensor s1 was placed near the exciter on the perpendicular surface, the second one (s2) on the opposite surface to the transmitter near the edge and the third sensor (s3) was on the same surface as the first one but close to the opposite edge. the generated signal can be in either form of continuous noise or harmonic signals [13]. the evaluation of the signals response to the transmitted pulse in the tested object can be performed in two basic systems of analysis. in the first case, the resulting shift in relation to time is evaluated, and in the second case, the signals are converted to the frequency domain and a frequency analysis is performed. the results used in this paper are from the time domain. subsequently, all specimens were provided by an initial notch and selected specimens were tested in a three-point bending configuration. selected specimens will be tested at the centre of excellence telč, czech republic in the four-point bending configuration, and x-ray computed tomography of these specimens is also planned – specimens are bellow marked ct. analysis of results of these fracture tests are being prepared and they are not the matter of this paper. the compressive strength value was determined at the parts of specimens obtained after the three-point bending fracture tests were performed. the auxiliary loading device with two (upper and lower) square steel auxiliary platens was placed in the hydraulic testing machine. the specimens were placed between the auxiliary plates in such a way that the load was applied perpendicularly to the compaction direction, as can be seen in figure 4. the test was carried out in accordance with the principles of standards čsn en 196-1 [14] and bs 1881: part 119 [15]. increasing load was applied continuously until the failure when the loading force decrease. the maximum load applied to each specimen was recorded. the informative compressive strength value of material of the specimen was calculated by dividing the maximum load by the area of contact – nominal surface area 40 × 20 mm2. 100 vol. 15/2018 characterization of cement-based composite figure 3. illustration of the ultrasonic pulse method experimental set-up (transmitter t, sensors s1, s2, and s3). figure 4. the compression test of specimens: the auxiliary loading device with two auxiliary platen was placed in the hydraulic testing machine with selected specimens prepared for the test (left); specimen after testing. 3. results all measured data, results of the ultrasonic pulse method and informative compressive strengths for test specimens are shown in table 1. basic statistics – mean values and standard deviations – of corrected selected data, as well as coefficients of correlation are introduced in tables 2 and 3. the ratio between maximum amplitudes r21 and r31 is called signal attenuation at the constant length – numbers 1, 2 and 3 correspond to the sensors s1, s2, and s3, index “b” or “a” means measurement before or after temperature loading. the outcome of applied procedure is a proportion of ratio r of temperature loaded specimen (r21,a or r31,a) to ratio r of the non-degraded specimen (r21,b or r31,b) – labelled as r21 and r31, respectively. 4. conclusions after the exposure of cement-based composites to elevated temperatures their mechanical parameters are reduced [4]. therefore, the presented paper was focused on the evaluation of the damage degree of finegrained cement-based composite specimens exposed 101 i. rozsypalová, m. vyhlídal, r. dvořák et al. acta polytechnica ctu proceedings specimen bulk density γ temperature t ratio r21,b ratio r31,b ratio r21,a ratio r31,a ratio r21 ratio r31 compressive strength fc,i kg·m−3 ◦c – – – – – – mpa 01-100 ct 1912 100 0.0140 0.0276 0.0406 0.0545 2.900 1.975 – 02-100 1875 100 0.0190 0.0527 0.0089 0.0146 0.468 0.278 4.61 03-100 1847 100 0.0651 0.0425 0.0074 0.0113 0.114 0.266 3.58 04-100 1906 100 0.0154 0.0154 0.0098 0.0153 0.640 0.993 8.21 25 1925 100 0.0949 0.1391 0.0578 0.0916 0.609 0.659 6.03 26 1958 100 0.3462 0.4216 0.1154 0.1450 0.333 0.344 – 27 2009 100 0.2550 0.4041 0.0833 0.0816 0.327 0.202 8.44 05-200 ct 1912 200 0.0120 0.0289 0.0026 0.0111 0.217 0.384 – 06-200 1988 200 0.0156 0.0206 0.0042 0.0121 0.270 0.587 6.91 07-200 1956 200 0.0262 0.0296 0.0081 0.0180 0.309 0.607 4.68 08-200 1844 200 0.0255 0.0216 0.0072 0.0071 0.282 0.327 7.73 09-400 ct 1958 400 0.0166 0.0200 0.0037 0.0025 0.223 0.125 – 10-400 1993 400 0.2805 0.4328 0.0375 0.0609 0.134 0.141 4.42 11-400 2020 400 0.1404 0.1638 0.0166 0.0239 0.118 0.146 4.29 12-400 1975 400 0.1546 0.1421 0.0232 0.0240 0.150 0.169 5.21 13-600 ct 1951 600 0.0792 0.0748 – – – – – 14-600 1972 600 0.0799 0.0974 – – – – 1.88 15-600 1997 600 0.1051 0.1930 – – – – – 16-600 2017 600 0.1448 0.1586 – – – – 3.93 17-800 ct 1974 800 0.1554 0.1405 – – – – – 18-800 1942 800 0.1017 0.2213 – – – – – 19-800 1887 800 0.2204 0.1781 – – – – – 20-800 1899 800 0.0895 0.1810 – – – – – 21-1000 ct 1925 1000 0.0866 0.0901 1.5438 1.1155 17.82 12.38 – 22-1000 1953 1000 0.1186 0.1631 0.3696 0.4316 3.118 2.647 1.96 23-1000 1958 1000 0.1208 0.1438 0.6280 0.9794 5.200 6.809 1.67 24-1000 1987 1000 0.1259 0.1394 0.3037 0.3264 2.413 2.341 1.45 table 1. measured data, results of the ultrasonic pulse method and informative compressive strengths. temperature t bulk density γ ratio r21 ratio r31 compressive strength fc,i ◦c kg·m−3 – – mpa 100 1919 (53) 0.475 (0.148) 0.457 (0.308) 6.17 (2.15) 200 1925 (63) 0.270 (0.039) 0.476 (0.142) 6.44 (1.58) 400 1986 (26) 0.156 (0.460) 0.145 (0.018) 4.64 (0.50) 600 1984 (29) – – 2.91 (1.45) 800 1925 (40) – – – 1000 1956 (25) 3.577 (1.449) 3.933 (2.496) 1.69 (0.26) table 2. mean values (standard deviations) of selected data. to high temperatures via ultrasonic pulse method. resulting data from this method are in quite high correlation with loading temperature (coefficient of correlation was in range 0.63 to 0.92), as well as with compressive strength of composite determined after the temperature loading (coefficient of correlation was in range –0.59 to –0.91). the very high correlation is between ratios r21 and r31 (coefficient of correlation was in range 0.97 to 1.00) which means that it is sufficient placed the sensors on the surface in the case of a large structure. the strong relationship between compressive strength and temperature for this special cement-based composite with spherical glass aggregates (coefficient of correlation was in range 102 vol. 15/2018 characterization of cement-based composite γ t r21 r31 fc,i bulk density γ 1.00 temperature t 0.23 | 0.29 1.00 symm. ratio r21 –0.06 | 0.14 0.63 | 0.92 1.00 ratio r31 –0.03 | 0.12 0.70 | 0.92 0.97 | 1.00 1.00 compressive strength fc,i –0.19 | –0.63 –0.80 | –0.97 –0.66 | –0.91 –0.59 | –0.90 1.00 table 3. coefficients of correlation (dimensionless) of selected data: all values (table 1) | mean values (table 2). –0.97 to –0.80) was confirmed in this study. the ultrasonic pulse method as a non-destructive method seems to be well-suited for determining the damage degree caused by elevated temperature in case of this special cement-based fine-grained composite. list of symbols γ bulk density before temperature loading at temperatures higher than 100 ◦c fc,i informative compressive strength r21,a the ratio of the maximum amplitude from sensor s2 to the maximum amplitude from sensor s1 after temperature loading at temperatures higher than 100 ◦c r21,b the ratio of the maximum amplitude from sensor s2 to the maximum amplitude from sensor s1 before temperature loading at temperatures higher than 100 ◦c r31,a the ratio of the maximum amplitude from sensor s3 to the maximum amplitude from sensor s1 after temperature loading at temperatures higher than 100 ◦c r31,b the ratio of the maximum amplitude from sensor s3 to the maximum amplitude from sensor s1 before temperature loading at temperatures higher than 100 ◦c r21 r21 = r21,a/r21,b r31 r31 = r31,a/r31,b t temperature acknowledgements this outcome has been achieved with the financial support of the czech science foundation under project no. 16-18702s “amiri – aggregate-matrix-interface related issues in silicate-based composites”. besides the funding project, the controlled temperature loading of the test specimen was carried out under the thorough supervision of technician pavel kropáček at institute of physics, faculty of civil engineering, brno university of technology. references [1] p.-c. aïtcin, s. mindess. sustainability of concrete. modern concrete technology. spon press, new york, 2011. [2] a. m. neville. properties of concrete. pearson education limited, essex, 2011. [3] i. hager. behaviour of cement concrete at high temperature. bulletin of the polish academy of science 61(1):145–154, 2013. [4] s. lim. effects of elevated temperature exposure on cement based composite materials (doctoral dissertation), 2015. [5] b. l. karihaloo. fracture mechanics and structural concrete. concrete design and construction. longman scientific & technical, new york, 1995. [6] s. kumar, s. v. barai. concrete fracture models and applications. springer, berlin, 2011. [7] j. farran. contribution mineralogique a l’etude de l’adherence entre les constituants hydrates des ciments et les materiaux enrobes. revue des materiaux de construction (490–491,492):155–157,191–209, 1956. [8] rilem. interfacial transition zone in concrete. in rilem state-of-the-art report. e & fn spon, london, 1996. [9] rilem. second international conference on the interfacial transition zone in cementitious composites. e & fn spon, new york, 1998. [10] k. l. scrivener, a. k. crumbie, p. laugesen. the interfacial transition zone (itz) between cement paste and aggregate in concrete. interface science 12(4):411–421, 2014. [11] g. prokopski, j. halbiniak. interfacial transition zone in cementitious materials. cem concr res 30(4):579–583, 2004. [12] j. p. ollivier, j. c. maso, b. bourdette. interfacial transition zone in concrete. advn cem bas mat 2(1):30–38, 1995. [13] n. j. carino. the impact-echo method: an overview. in 2001 structures congress & exposition. 2001. doi:10.1061/40558(2001)15. [14] čsn en 196-1. 2016 methods of testing cement part 1: determination of strength. [15] b. 1881. testing concrete part 119 method for determination of compressive strength using portions of beams broken in flexure (equivalent cubemethod), 1983. 103 http://dx.doi.org/10.1061/40558(2001)15 acta polytechnica ctu proceedings 15:99–103, 2018 1 introduction 2 experimental programme – materials, specimens, methods 3 results 4 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0079 acta polytechnica ctu proceedings 12:79–82, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app automatic generation of road infrastructure in 3d for vehicle simulators adam orlický faculty of transportation sciences, czech technical university in prague, konviktská 20, prague 1, czech republic correspondence: orlicada@fd.cvut.cz abstract. one of the modern methods of testing new systems and interfaces in vehicles is testing in a vehicle simulator. providing quality models of virtual scenes is one of tasks for driver-car interaction interface simulation. nowadays, there exist many programs for creating 3d models of road infrastructures, but most of these programs are very expensive or canâťt export models for the following use. therefore, a plug-in has been developed at the faculty of transportation sciences in prague. it can generate road infrastructure by czech standard for designing roads (čsn 73 6101). the uniqueness of this plug-in is that it is the first tool for generating road infrastructure in nurbs representation. this type of representation brings more exact models and allows to optimize transfer for creating quality models for vehicle simulators. the scenes created by this plug-in were tested on vehicle simulators. the results have shown that with newly created scenes drivers had a much better feeling in comparison to previous scenes. keywords: vehicle simulator, road design, 3d modeling. 1. introduction with emerging technologies, many branches of the technical sciences have appeared during the last few decades. one of such branches is computer graphics, whose one of the sections is the 3d modeling. it deals with displaying the real world in virtual threedimensional space. for this the 3-d modeling is used. in road transportation 3d modeling is used for the planning and design of newly projected or reconstructed roads and highways. generation of 3d models in special software is usually restricted by local laws, besides, the price of the software is rather high. most of these applications allow to create project documentation for the designed roadways, display the developed models in 3d and, in the case of highly sophisticated software, create visualization and animation. typically, exporting is not really possible, that is why this software can’t be used for other purposes[1]. in transportation, virtual reality is widely used in creation of interactive vehicle simulators. the real world can be displayed in virtual scenes with the help of projectors, or on monitors. one of the advantages of exploitation of driving simulators is a possibility to evaluate driver behavior in the conditions and situations that are precisely set up in driving scenarios. it is important that while driving the simulator a driver would get a feeling maximally approaching the experience of real driving. quality 3d models can contribute to the improvement of the visual perception. in most modern driver simulators, the road infrastructure isn’t generated according to these standards. all of this provides a motivation for the creation of a tool (as a plug-in for 3d modeling program) which is able to generate 3d road infrastructures according to czech standard[2], developed for a minimal price and for wide-ranging applications. for this purpose, we were looking for a software, where both design drawing and 3d modeling are possible. besides the generation of road infrastructure models, a tool (plug-in) should be able to complete a scene around roads in such a way, that the resulting virtual scenery would be as close to reality as possible. 1.1. driving simulator scene requirements there are rather specific requirements for 3d models of virtual scenes. driving simulator scenarios are most commonly created using the tin (triangulated irregular network) model[3]. with such a representation, the models are being created within an irregular network of triangles. these models should be as maximally close to reality as well as being simple (composed of the minimum possible number of triangles). another important requirement are the models, which are not visible to a driver (proband), needn’t be created. the last requirement would be the complexity of a model inversely proportional to the object’s distance from the road axis. application of these approaches makes it possible to use the less powerful hardware for creating virtual sceneries for driving simulator. optimized 3d scene models allow higher frame rates for the projected scenes, which provides a better illusion of a real world. 79 http://dx.doi.org/10.14311/app.2017.12.0079 http://ojs.cvut.cz/ojs/index.php/app adam orlický acta polytechnica ctu proceedings figure 1. tin models. 2. plug-in road creator for rhino plug-in road creator for rhino was created for 3d modeling and visualization program rhinoceros. the program is focused on creating models in nurbs (non-uniform rational basis spline) representation, where objects are mathematically described with the help of control points and their importance[4]. thus, the objects are more precise and don’t contain sharp edges. this kind of representation is also valuable for the possibility of further transformation of 3d model into tin model, for which it is possible to optimize the count of triangles of the model itself and thus optimize models for a driving simulator scene. figure 2. nurbs models. in the created plug-in it is possible to generate 3d models of road infrastructure (according to czech standard čsn 73 6101) into the rhinoceros software[2]. the standard is about the creation of designing roads and highways in the czech republic. other important elements in the infrastructure are generated per the corresponding standards and technical specifications. the plug-in was designed in a way that allows fast generating of the driver simulator scenes. that’s why databases with modelled objects was created for plugin. overall two databases have been formed. the first one is the database named "trees" that contains models of trees, bushes and grass and it is necessary for the completion of vegetation in the scene. the second database named "buildings" contains building and specific elements present around the roads. plug-in contains 54 new functions, due to which the scene creation is not as difficult as it used to be before. for user convenience, all the functions have been placed on a specially created control panel. for possible viewing of the scene model, the plug-in structures have newly created objects as layers. 2.1. plug-in functions the plug-in was created for the creation of driving scenarios. the creation of road infrastructure deploys five groups of functions of the plug-in. those include: • 2d road, • 3d road, • extras, • nature, • city. the functions can be accessed from the interface toolbar (see figure 3 for reference). figure 3. plug-in functions (interface toolbar). the process itself consists of four sequential stages. schematic and logical representation is shown in figure 11. stage 1 (model pre-processing) serves for the creation of a terrain model which is the basis of a road network. for transferring the 3d terrain model into the contour lines the function "contour lines" from the first group is applied (see figure 4). figure 4. terrain transfer to contour lines. at stage 2 basic curves for the future road communication are created using the functions from 2d group. with the help of horizontal (successive use of functions "tangential polygon", "arch with transition curve", and "vertical alignment" + "tangential polygon" from road 2d group) and vertical (functions "vertical alignment" and "parabolic arch") alignment combination a road track is created (see figure 5). 80 vol. 12/2017 automatic generation of road infrastructure in 3d by adding the cross-section type from the defined database of the cross-sections one defines the character of the road being created, i.e. number of lanes, their width etc. the database contains the crosssections corresponding to classes of the road defined in czech standard čsn 73 6101. figure 5. creation of road track. the plug-in provides automatic modification of the cross-sections according the curvatures and with respect to the location coordinates (e.g., at intersections) on the created road track (see figure 6). figure 6. basic road curves. now, when the basic curves are created, the 3d model of the road track can be created by using the function "3d road" from the "road 3d" menu section – the 3rd stage of the process. for the integration of the model into the terrain the function "slopes" is applied. see the result in figure 7. figure 7. road with slopes, 3d model. the 4th stage is the final one, it results in a complete road track model that includes such attributes as barriers, traffic signs, pillars etc. see example on figure 8. figure 8. complete road track model. with the help of the plug-in described, it is also possible to create other scenario attributes, such as nature, buildings etc. see the final model example in figure 9. figure 9. final 3d model. for the creation of the basic city infrastructure, a group of functions "city" shall be applied. these functions allow creation of different intersection types (see example in figure 10), sidewalks, pedestrian crossings etc. figure 10. city infrastructure – roundabout model. 3. conclusions the developed plug-in is nowadays unique regarding its possibility to create nurbs models of road infrastructure according to standard čsn 73 6101. unlike the tin models, the nurbs models are more precise, while there is a possibility of transformation of nurbs models into tin models with setting the quality of created tin models in rhinoceros software. the advantage is quality of tin models can be optimized for driving simulators. during the transferring 81 adam orlický acta polytechnica ctu proceedings figure 11. road infrastructure creation process. it has been accomplished that the models in simulator are not containing visible sharp transitions and are being formed with minimal possible number of triangles. the created 3d road infrastructure models can be also used for visualization, the plug-in can also be applied for schematic drawings of track situation and longitudinal profile, which can be used for the project documentation. besides, the major advantage of this plug-in is that it significantly reduces the time spent at modeling. creation of 3d scenery models with road infrastructure in this plug-in is the matter of several minutes. after processing of the created infrastructure in the 3d modeling engine the scenario for simulation is ready (see figure 12). figure 12. final simulation scenario. the plug-in is presently being used at transportation faculty at czech technical university in prague as a main tool for generation of scenes for driving simulators. the functions of these plug-ins were also used for the creation of animations and visualizations in the final theses of students studying at the faculty of the author. references [1] road design by autocad civil 3d and 3d max design. youtube [online]. 2014 [cit. 2016-04-29]. dostupné z: https://www.youtube.com/watch?v=eeo0wabu7ow. [2] čsn 73 6101 projektování silnic a dálnic. praha: český normalizační institut, 2004. [3] počítačová grafika: základy počítačové grafiky. ict kompetence [online]. olomouc: katedry technické a informační výchovy pdf up. [cit. 2016-04-29]. dostupné z: http://www.kteiv.upol.cz/frvs/ictkubricky/?page=pocitacova-grafika/pocitacova-grafika. [4] rhino developer docs [online]. robert mcneel & associate, ©1997-2016 [cit. 2016-04-30]. dostupné z: http://developer.rhino3d.com/. 82 acta polytechnica ctu proceedings 12:79–82, 2017 1 introduction 1.1 driving simulator scene requirements 2 plug-in road creator for rhino 2.1 plug-in functions 3 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.10.0061 acta polytechnica ctu proceedings 10:61–64, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app stabilization and specification of characteristic value of geotechnical parameters ivan vaníček czech technical university in prague, faculty of civil engineering, thákurova 7, prague 6, czech republic correspondence: ivan.vanicek@cvut.cz abstract. soil stabilization, process, by which the soil properties are improved, is getting growing attention during last periods. paper is focused on application for earth structures. methods of soil stabilization together with their applicability for different soils are discussed firstly. the sensitivity of the process of the cautious estimation of the characteristic values of geotechnical parameters which are subsequently used for the design, is shown on the base of the logical scheme of the geotechnical design report for earth structures. finally the need of the laboratory tests for the closer specification of the soil stabilization technology and also for the determination of the geotechnical data are discussed. keywords: stabilization, earth structure, eurocode 7, characteristic value. 1. introduction soil stabilization is one of the techniques of soil improvement. together with soil reinforcement it is the most applied techniques. soil improvement methods can be divided with respect to the characteristic values specification into two basic subgroups [1]: • diffusive soil improvement; • discrete soil improvement. in the first case, typical for soil stabilization, the design counts with new properties for geological environment in general, better speaking for individual quasi-homogeneous layers of the environment. in the second case soil properties remain the same (on the previous values), as well the parameters of the reinforcing element (most often geosynthetics). the behaviour of the composite therefore depends on the interaction of these two basic materials and limit states of failures should be determined also for this reinforcing element. for earth structures the improvement of individual soil layers of fill or surface of ground (subsoil) has dominant application role. it is typical for earth structures of transport engineering including airports. however it is connected also with subsoil improvement of large halls, especially when they are founded partly in cut and partly on fill. soil stabilization, mixing soil particles with different additives, can be realized either on site or in mixing centre situated of-site with subsequent transport on the place of application. additives are different, either according to the soil type or with respect of certain demands. the following additives can be applied [2]: • lime, cement, plaster; • industrial by products as fly ash, energy-gypsum, slag (cinder) [3]; • additional soil mechanically stabilized soil; • several chemical additives (resins, bitumen). the paper is for the simplicity focused on the first possibility, on the application of lime or cement, whereas cement is typical for coarse soil and lime for fine grained soils. the recommendation in this direction are often connected with index plasticity. for cement application, the higher strength is reached, however the strength is rising up relatively slowly. the final result is more typical for road base or subbase or for layers just below floor (e.g. for above mentioned halls), where higher strength is typically demanded. subsequently more attention is focused on lime application, which is more suitable for fine soils and where quicker effect of application is needed. 2. lime stabilization quick lime is preferred for soil stabilization as it hydrates with water and calcium hydroxide develops. this reaction is exothermic. heat is created as the result of quick lime reaction with wet soil having quick impact on soil compaction which is after that for dryer soil better. therefore generally two impacts are distinguished, short term (having positive impact on soil compaction) and long term (having positive impact on strength increase). while short term effect occurs for any quick lime application, long term effect need a certain percentage of quick lime. broadly speaking this boundary is close to 2 % of added lime to dry soil. particle flocculation starts for higher content of lime, changing grain size distribution curve with lower content of fine particles. soil creates clods and change the structure. therefore for fine grain soils the stabilization process with lime leads to: • better soil workability; • strength increase. from this point of view the long term effect is more important for increase of characteristic values of geotechnical parameters, however requires higher lime 61 http://dx.doi.org/10.14311/app.2017.10.0061 http://ojs.cvut.cz/ojs/index.php/app ivan vaníček acta polytechnica ctu proceedings content, usually in the range of 2 to 6 %. stabilization effect connected with long term effect is therefore connected with reaction of lime with clay particles, subsequently leading to: • growth in strength and the subsoil bearing capacity in general due to the creation of cementing bonds resulting from soil-lime pozzolanic reaction, which, however, develop in time and therefore strength tests are performed with a time lag of mostly 28 days; • reduction in the susceptibility to swelling and shrinkage; • reduction in moisture content is caused both by absorption, extraction of moisture from the surrounding soil after the lime type is changed into slake lime, and by hydration heat as was mentioned above. at the same time, however, the new reaction of lime with clay minerals leads to a change in plasticity thus affecting another factor relevant for workability. • two additional notes are needed [4]; • for higher content of quick lime (above 4-6 %) there is no another reduction of plasticity; • for the development of cementing bonds clay particles are needed, at least 10 %. 3. technology effect the application of lime stabilization on site prevails at the present days. after areal spreading of the demanded content of lime the soil and lime are mixed with the help of cutter or another machineries to the required depth. such modified layer is levelled by grader or bulldozer and subsequently compacted by roller. a great attention is devoted to the moisture content at the time of compaction, to guarantee that soil after compaction will not have more than 5 percent of air in pores. therefore some specialists devote a great attention to the mellowing maturing, to the time between soil and lime mixing and final compaction. therefore after mixing only light compaction is recommended and after about 24 to 72 hours, followed by second mixing supplemented by additional increase of water the final compaction is applied. it is consistent that this recommendation has some limits, as for example for improvement of railway tract subsoil during rail heave [5]. from this technology point of view the application of lime stabilization is not recommended during late autumn, as stabilized soil can freeze. for such case addition of cement is recommended after short delay after lime. nevertheless this recommendation has general character, as lime can prepare soil for cement application. this effect can be reached by some special mixed additives, e.g. for the czech republic it is dorosol, which contains a certain percentage of lime and the rest are hydraulic ingredients. in this case a special attention is devoted to dusting lowering. 4. logical scheme of the earth structure design contemporary version of the eurocode 7 geotechnical design contains four important steps of the geotechnical structure design and performance as specification of geological model, geotechnical model, calculation model and ending with structure realization, performance connected with structure control. these four phases can be used for the specification of the risk with which the design and performance is connected, first of all from the point of view of uncertainties and indeterminateness. for earth structures this scheme is presented in fig. 1 in the form of the geotechnical design report gdr. geological model should be specified not only for the ground but also for the borrow pit, from which fill will be constructed. geotechnical investigation report or ground investigation report gir includes not only geological model but also results of the laboratory and field tests of the ground and theirs presentation in the form of measured or derived geotechnical parameters. from the borrow pit technological samples are taken and subsequently tested in the laboratory. firstly it is compaction test most often proctor standard test. range of the geotechnical investigation as well the range of performed tests is connected with risk, with which the design and performance is connected. eurocode 7 recommends three degrees for this specification with the help of so called geotechnical categories (gc), when into: • 1 gc belongs simple structures in modest geological and geotechnical conditions. these structures are designed with the help of up to day experiences calculation of limit states is not needed. generally into first gc belongs fills with high up to 3 m with good ground conditions. however this is not the case for the ground of fill which need improvement soil stabilization. • 3 gc belongs very demanding structures, or structures with complicated ground from the geological and geotechnical point of view or structures connected with extreme risk for surrounding environment in the case of failure. • 2 gc belongs all other structures. for the structures falling into 3 gc it is necessary always determine mechanical-physical properties either for ground or for fill. for material of fill the test are performed on samples compacted in laboratory and for which moisture content and dry density are falling into the accepted (recommended) range. the same procedure is recommended for structure which belongs to the second gc, however with one exception. in the case of 2 gc connected with lower risk the mechanical-physical properties can be selected from 62 vol. 10/2017 specification of characteristic value of geotechnical parameters figure 1. gdr scheme for earth structures. standard tables, based on the correlation of index properties and mechanical-physical properties. such correlation for compacted fill is e.g. used in codes for small dams, (in the cr it is e.g. code čsn 75 2410 small water reservoirs). however for stabilized soils, firstly where the strength increase is needed, such tables can be hardly used, as there is a shortage or correlations, and therefore lab tests are needed. in this case very good cooperation between project designer and specialist on the lab tests is preferred, namely with respect to the tests performed on samples which were prepared by the nearly same way as is proposed stabilization on the site. 5. sensitivity of the characteristic values determination as stated already lab tests of the mechanical-physical properties are needed for structures belonging into 3 gc and 2 gc with moderate and high risk. after that the results are statistically evaluated and after subsequently they are used for the determination of the characteristic values of the geotechnical parameters. here ec 7 defines the characteristic values as “selected as a cautious estimate of the value affecting the occurrence of the limit state”. this expression should be taken very carefully, as: • “selected” emphasizes the engineering judgement; • “cautious estimate” emphasizes that a certain conservatism should be applied; • “limit state” emphasizes that selected values should have close relation to the given limit state. nevertheless also other aspects are playing significant role during the process of selection: • amount of information and a degree of confidence to these information; • volume of ground affected by expected limit state of failure (e.g. slip surface) or an ability of the structure to transfer loads when both aspects are connected with an ability to compensate weaker zones by stronger ones [6]. therefore for slip surface of failure with significant length it is possible to select characteristic values closer to the average ones, while for shorter slip surface the cautious estimate will be more conservative. for all that it can be specified that the characteristic values selection is one of the most sensitive task of the project designer. probability of verification of the selected values is very low. quality of ground is usually not controlled, only of the fill. but even for the fill the control is going mostly via quality of compaction, whether the moisture content and dry density are falling into prescribed range. direct tests of mechanical-physical properties are rather exceptional, e.g. permeability control for sealing system of landfills. 63 ivan vaníček acta polytechnica ctu proceedings 6. laboratory measurement laboratory tests are needed for soil stabilisation where the main aim is strength increase. optimal content of the additives should be determined as well with mechanical-physical properties, from which the designer selects characteristic values presented some interesting results performed within her diploma theses, devoted to the additives research [7]. for given soil, characterized as clay with low plasticity (cl) she concluded, that it is material unsuitable or conditionally suitable which is also sensitive to freezing. she tested stabilization by quick lime, cement and dorosol, and she focused on: • impact of lime on change of consistency limits; • impact of additives amount on increase of bearing capacity (controlled via cbr test); • impact of additives amount and time on the resistivity of stabilized soils to freezing cycles. follow-up diploma and phd thesis are focused not only on classical soils, but also on alternative aggregates as flying ash, when different additives are tested [3]. last attention is devoted to the combination of additives with randomly distributed short synthetic fibres. the evaluation of these tests is mostly via possibility of application of stabilized soils into road foundation road base or sub-base layers, e.g. according to the czech code čsn 73 6125. up to now lower attention is devoted to the additives application for the own fill of the earth structure. there the properties as shear parameters and deformation modulus have more informative impact, as they are needed for the solution of the limit states uls ultimate limit state or sls serviceability limit state. the problem is very sensitive, as a stiffness increase is function of time and the designer should solve the limit states not only for short term conditions, but also for long term conditions. 7. conclusion the application of different additives for improvement of soil properties has growing tendency. from the view of lime application on fine grained soils it is necessary to distinguished between improvement from the workability view with short term effect and improvement with long term effect, when strength increase has permanent character. in the course of additives application for earth structures of transport engineering the designer should firstly specify his/her intention, whether the process of improvement is focussed on the ground and own fill or on the construction layers, as the applications (and conditions) are different. time effect is playing important role as properties are time dependant, mostly from the view of long term bonds having decisive impact on the mechanical-physical properties. therefore the lab tests are needed pro for demanding structures with moderate and high risk and should be focused not only on the specification of additive amount but also on the determination of soil properties reported as characteristic values. from the view of safety the designer of earth structures has to face many uncertainties with respect to the soil properties, (compared to the designer dealing with man-made construction materials) as subsequent control of the compacted soil is limited and therefore the transfer of experiences from one site to the another one is limited. acknowledgements the paper was prepared with financial support from the research project te0 120168 of the technology agency of the cr: cesti centre of effective and sustainable transport infrastructure. references [1] i. vaníček. significance of geotechnical investigation for geotechnical structure design. in proceedings of 43rd conference with international participation: foundation engineering. brno, cgts cice, pp. 27–40. brno, 2015. [2] m. vaníček, i. vaníček. earth structures in transport, water and environmental engineering. springer, dordrecht, 2008. [3] v. mráz. research of stabilized fly ash properties and its application for earth structures of transport engineering (in czech). ph.d. thesis, czech technical university in prague, faculty of civil engineering, 2016. [4] h. m. greaves. an introduction to lime stabilization. in lime stabilisation. proceedings of the seminar held at loughborough university, civil ans building engineering dept on 25 september 1996. [5] j. h. smith. construction of lime or lime plus cement stabilized cohesive soils. in lime stabilisation. proceedings of the seminar held at loughborough university, civil ans building engineering dept on 25 september 1996. [6] i. vaníček. jubilee volume: andreas anagnostopoulos 50 years of service at the national technical university of athens, chap. earth structures sensitivity of the design to the characteristic geotechnical parameters values determination., pp. 543–552. tsotras, 2015. [7] vrbová. soil stabilization. master’s thesis, czech technical universiti in prague, faculty of civil engineering, 2008. 64 acta polytechnica ctu proceedings 10:61–64, 2017 1 introduction 2 lime stabilization 3 technology effect 4 logical scheme of the earth structure design 5 sensitivity of the characteristic values determination 6 laboratory measurement 7 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0085 acta polytechnica ctu proceedings 7:85–90, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app modeling of anisotropic growth and residual stresses in arterial walls anna zahna, daniel balzania, b, ∗ a technische universität dresden, institute of mechanics and shell structures, 01062 dresden, germany b dresden center for computational materials science (dcms), 01062 dresden, germany ∗ corresponding author: daniel.balzani@tu-dresden.de abstract. based on the multiplicative decomposition of the deformation gradient, a local formulation for anisotropic growth in soft biological tissues is formulated by connecting the growth tensor to the main anisotropy directions. in combination with an anisotropic driving force, the model enables an effective stress reduction due to growth-induced residual stresses. a method for the imitation of opening angle experiments in numerically simulated arterial segments, visualizing the deformations related to residual stresses, is presented and illustrated in a numerical example. keywords: biomechanics, multiplicative anisotropic growth, residual stresses, opening angle. 1. introduction as living biological medium, arterial tissue is known to adapt to changes in its mechanobiological environment in order to improve the load-bearing capacities. such adaptation processes originate from growth and remodeling and may be indicated by the existence of residual stresses under externally load-free conditions. those residual stresses can be uncovered by extracting an arterial segment from the body and slicing it in radial direction. typically, the segment contracts in axial direction and springs open to a degree which depends on various factors, for example the height of the blood pressure [1]. being associated with a reduction of stress gradients in loaded states, residual stresses need to be included in numerical simulations to obtain a predictive stress distribution. computational methods to reduce stress gradients by incorporating residual stresses are proposed in the literature [2–5]. these methods are however rather engineering approaches, as no biological processes are included to obtain the residual stresses. focusing on growth as part of the adaptation processes, the addition of material, resulting in increasing wall thicknesses and diameters, is related to a reallocation of the stresses which may therefore allow for an optimized load bearing behavior. for the purpose of modeling these effects, an anisotropic growth model based on the multiplicative decomposition of the deformation gradient into a growth tensor and a remaining elastic part is presented. in this context, different approaches for anisotropic growth in soft biological tissues have been published, see e. g. [6, 7], where the anisotropy of the growth mechanism is based on the definition of macrostructural growth directions as for instance the radial direction. although the anisotropy of the material is taken into account, isotropic strainor stress-based quantities are used as driving forces in the evolution equations characterizing the growth tensor. decomposing the growth tensor itself into two parts related to the main material anisotropy directions, the presented growth model establishes a relation between the growth tensor and the fibrous tissue structure, which yields a local formulation and therewith supersedes a geometry-specific definition of the growth directions. aiming at rendering the growth and stress reduction mechanism more effective, an anisotropic stress-based driving force is formulated. in addition, a procedure for the calculation of the deformations related to growth-induced residual stresses in numerically simulated arterial walls is developed. the entire framework is illustrated with the aid of a numerical example on an idealized arterial segment. 2. computational framework having been introduced in the context of biomechanics by rodriguez et al. [8] in 1994, the multiplicative decomposition f = fefg of the deformation gradient f into a growth part fg and an elastic part fe is nowadays a common approach for the description of growth in soft biological tissues. the growth deformation is assumed to be stress-free, such that stresses arise only from the elastic part ce = f −tg cf −1g of the strains. utilizing constitutive equations for hyperelastic materials, the stress tensor se in the intermediate configuration can be computed from se = 2 ∂ψ∂ce . here, the formulation for fiber-reinforced soft biological tissues by balzani et al. [9] is applied, whereby polyconvexity of the strain energy function and thus ellipticity and existence of minimizers are ensured. the growth behavior is governed by the specific form of the growth tensor, which involves one or several local variables referred to as growth factors ϑi. the evolution equations for these growth factors ϑ̇i = kϑ(ϑi) φi(zi), are formulated as function of the growth-driving quantity φ(zi) and the growth function kϑ including the time-dependency. 85 http://dx.doi.org/10.14311/app.2017.7.0085 http://ojs.cvut.cz/ojs/index.php/app anna zahn, daniel balzani acta polytechnica ctu proceedings in this section, a model for anisotropic growth, relating the growth tensor fg to the anisotropy directions of the fibrous tissue material, is presented in subsection 2.1. applying this framework to idealized arterial segments under internal pressure, residual stress distributions that qualitatively reflect experimental observations can be generated. in subsection 2.2, the focus is directed on the numerical calculation of the deformations which arise from resecting and slicing previously simulated, residually stressed arterial segments. 2.1. model of anisotropic growth related to two fiber families an effective approach to reduce stresses is the enlargement of the cross-sectional area. accounting for this, the basic form fg = ϑi + (1 −ϑ) a⊗a, describing a deformation with factor ϑ in all directions perpendicular to the direction given by the vector a, is chosen among different existing approaches for the growth tensor, see e. g. [10]. in constitutive material models for arterial tissues, two main fiber directions a(a) with a = 1, 2 are usually considered to approximate the complex layer-wise arrangement of collagen fibers. assuming the anisotropy of growth to be linked with the fibrous tissue composition, a multiplicative decomposition of the growth tensor fg = f (ii)g f (i) g (1) into two separated parts f (a)g = ϑ (a)i + ( 1 −ϑ(a) ) m(a) with (2) m(a) = a(a)g ⊗a (a) g , a = i,ii (3) is considered in order to account for two different anisotropy directions defined by the vectors a(i)g and a (ii) g . this formulation is based on the assumption that in arterial walls, even in patient-specific ones, a stress state exists with one negative principle stress showing in radial direction and two positive principle stresses in a plane of tensile loading (which coincides with the axial-circumferential plane in perfect-tube arteries). this can be supposed to be the reason for the observation that mainly two collagen fiber families, oriented in directions within this plane, occur in arterial walls. the resulting tissue structure might enable an optimized load-bearing mechanism, cf. [11, 12]. in more detail, these two fiber families may favorably be oriented symmetrically with respect to the direction of the highest principal stresses and thus, growth perpendicular to these directions would be desired by the tissue to reduce the stresses. consequently, the growth orientation vectors a(a)g are defined as angle bisectors of the fiber direction vectors according to a(i)g = a(1) + a(2) ||a(1) + a(2)|| and (4) a(ii)g = a(1) −a(2) ||a(1) −a(2)|| . (5) for the determination of the growth factors ϑ(a), two evolution equations ϑ̇(a) = kϑ(ϑ(a)) φ(a)(z(a)), a = i,ii (6) have to be considered. therein, the function kϑ(ϑ(a)) =   k+ϑ [ ϑ+ −ϑ(a) ϑ+ − 1 ]m+ ϑ , φ(a) > 0 k−ϑ [ ϑ(a) −ϑ− 1 −ϑ− ]m− ϑ , φ(a) < 0, (7) which has been adopted from lubarda & hoger [13], scales the growth velocity by a factor k+/−ϑ and restricts the range of the growth factors to [ϑ−,ϑ+]. the function φ(a)(z(a)) defines the quantity which induces growth. recent publications examining stressdriven growth [6, 7, 14] propagate the utilization of the trace of the elastic part of the mandel stress given by φ(cese) = cese : i as measure for the pressure acting on the arterial wall. since this quantity is isotropic, it appears not to be adequate in combination with the anisotropic growth tensor and is therefore modified as follows. using the projections of the elastic part of the mandel stress into the anisotropy directions a(a)g , φ(a)(cese) = cese : m(a)e (8) with m(a)e = a (a) g ⊗a (a) g , (9) it can be ensured that growth in the plane perpendicular to a(a)g is only emerging if the stress in this direction is nonzero. the index “e” indicates, that the structural tensors m(a)e refer to the intermediate configuration. however, being eigenvectors of the growth tensor, the growth orientation vectors a(a)g do not change their direction during the transformation to the intermediate configuration, thus m(a)e = m(a). involving strain and stress quantites in the intermediate configuration, i. e. ce = f −tg cf −1g and se = 2 ∂ψ∂ce , the functions φ (a) depend on both growth factors ϑ(a), yielding a coupled system of evolution equations (6). using an euler backward scheme as for instance done in [7], the approximations ϑ̇(a) ≈ ϑ (a) n+1 −ϑ (a) n ∆t , a = i,ii (10) are adopted in order to compute the growth factors belonging to the actual time step n + 1. note that the index n + 1 is skipped from now on and only 86 vol. 7/2017 modeling of anisotropic growth and residual stresses in arterial walls quantities at the last time step are denoted by the index n. consequently, the local residuals r(i) := −ϑ(i) + ϑ(i)n + kϑ(ϑ (i)) φ(i)(ϑ(i),ϑ(ii)) ∆ t, r(ii) := −ϑ(ii) + ϑ(ii)n + kϑ(ϑ (ii)) φ(ii)(ϑ(i),ϑ(ii)) ∆ t have to fulfill r(i) = r(ii) = 0. to solve these nonlinear equations, a local newton iteration, characterized by index k, is performed. the linearized residuals linr(i) = r(i) ∣∣ ϑ(a),k + ∂r(i) ∂ϑ(i) ∣∣∣∣∣ ϑ(a),k ∆ϑ(i),k+1 + . . . . . . ∂r(i) ∂ϑ(ii) ∣∣∣∣∣ ϑ(a),k ∆ϑ(ii),k+1 = 0, linr(ii) = r(ii) ∣∣ ϑ(a),k + ∂r(ii) ∂ϑ(i) ∣∣∣∣∣ ϑ(a),k ∆ϑ(i),k+1 +. . . . . . ∂r(ii) ∂ϑ(ii) ∣∣∣∣∣ ϑ(a),k ∆ϑ(ii),k+1 = 0 are solved for the incremental updates ∆ϑ(i) and ∆ϑ(ii), and ϑ(i) and ϑ(ii) are updated until an appropriate abort criterion is fulfilled. during this iteration, the incremental updates are calculated by ∆ϑ(i),k+1 = r(i) a (i,i) ϑ ∣∣∣∣∣ ϑ(a),k − a (i,ii) ϑ a (i,i) ϑ ∣∣∣∣∣ ϑ(a),k ∆ϑ(ii),k+1, (11) ∆ϑ(ii),k+1 = a (i,i) ϑ r (ii) −a(ii,i)ϑ r (i) a (i,i) ϑ a (ii,ii) ϑ −a (i,ii) ϑ a (ii,i) ϑ ∣∣∣∣∣ ϑ(a),k , (12) where the abbreviations a (i,i) ϑ = − ∂r(i) ∂ϑ(i) , a (i,ii) ϑ = − ∂r(i) ∂ϑ(ii) , a (ii,i) ϑ = − ∂r(ii) ∂ϑ(i) , a (ii,ii) ϑ = − ∂r(ii) ∂ϑ(ii) are used. thus, the derivatives of the functions kϑ and φ(a) with respect to the growth factors have to be provided. differentiating eqs. (7) and (8) leads to ∂kϑ(ϑ(a)) ∂ϑ(a) =   m+ϑ ϑ(a) −ϑ+ kϑ(ϑ(a)), φ(a) > 0 m−ϑ ϑ(a) −ϑ− kϑ(ϑ(a)), φ(a) < 0 (13) and ∂φ(i) ∂ϑ(j) = ∂cese ∂ϑ(j) : m(i)e = ( ∂ce ∂ϑ(j) se + ce ∂se ∂ϑ(j) ) : m(i)e (14) = ( ∂ce ∂ϑ(j) se + ce ( 1 2 ĉe : ∂ce ∂ϑ(j) )) : m(i)e with i = i,ii and j = i,ii. herein, ĉe = 2 ∂se ∂ce is the elastic part of the tangent moduli in the intermediate configuration. the remaining unknown derivative of the strain ce with respect to the growth factor ϑ(a) has to be computed according to the chain rule by ∂ce ∂ϑ(a) = ∂ce ∂fg : ∂fg ∂ϑ(a) = ∂ce ∂fg : ∂fg ∂f (a) g : ∂f (a) g ∂ϑ(a) . (15) the first part is given through ∂ce ∂fg = −f −tg � ce −ce � f −t g (16) and the relations ∂fg ∂f (i) g = f (ii)g � i, (17) ∂fg ∂f (ii) g = i � (f (i)g ) t , (18) ∂f (a) g ∂ϑ(a) = i −m(a), a = i,ii (19) can be derived for the other parts.1 combining eqs. (17)/(18) and (19), the results ∂fg ∂ϑ(i) = f (ii)g −f (ii) g m (i), (21) ∂fg ∂ϑ(ii) = f (i)g −m (ii)f (i)g (22) are obtained and all required parts for the incremental updates of the growth factors in eqs. (11)–(12) are determined. given the final values of ϑi and ϑii, the 2nd piolakirchhoff stress s = f −1g sef −tg can be identified performing a pull back of se to the reference configuration. 2.2. visualization of residual stresses anticipating the results of the numerical example in section 3, where an idealized arterial segment under constant internal pressure is examined using finite element calculations, the question arises, how to determine the deformations which are related to residual stress states. in order to permit these deformations, a change of the boundary conditions is required, for instance the release of the right edge of a residually stressed arterial segment as shown in figure 1, which enables the opening of the segment by the angle ϕ. dealing with highly nonlinear material behavior and large deformations, this final deformation state can not be determined in one calculation step, thus a special procedure has to be applied. here, this procedure is proposed as follows. 1the symbol � is defined such that the relation (a � b)︸ ︷︷ ︸ =:d : (a ⊗ b) = aa ⊗ bb, dijkl = aik bjl (20) is valid, which for instance results in the calculation rule abc = ( a � ct ) : b. 87 anna zahn, daniel balzani acta polytechnica ctu proceedings ϕ ϕ 2 figure 1. definition of the opening angle ϕ in an idealized, symmetric arterial geometry. (1.) define the initial boundary value problem including the discretization of the unstressed configuration, apply the loads stepwise and hold them constant until a growth equilibrium state is reached. (2.) deactivate growth and remove the loads stepwise to determine the residually stressed state. due to growth, the obtained geometry differs from the initial reference configuration. (3.) save this grown, unloaded configuration by storing the nodal coordinates as well as the fiber orientation vectors and the residual cauchy stress σres := (det f )−1fsft at the integration points. (4.) use the saved nodal coordinates and fiber orientation vectors to define a new boundary value problem with boundary conditions that allow for the expected deformation. (5.) restore σres at each integration point and keep it throughout the following calculation. (6.) apply the residual stresses internally by updating the stress tensor according to s ← s + floadsres, where fload is a load factor ranging from 0 to 1, while growth is deactivated. in each step, the stress sres = det[f ]f −1σresf −t has to be computed corresponding to the actual deformation state. following this procedure, the wanted deformation state will arise in absence of external loads. 3. numerical example in a previous comparative study [15], an arterial segment with internal pressure and an axial displacement of 15 % of the initial length was considered for a comparison of two different forms of the growth tensor and two different driving forces. each of the approaches was able to provoke growth-induced circumferential stress smoothening, but the approach presented in subsection 2.1 turned out to be the most effective, generating the most beneficial stress distributions at the lowest growth requirement, measured by the determinant of the growth tensor. therefore, a similar boundary value problem making use of this approach will be analyzed here in more detail, focussing in particular on the simulation of the opening artery. the half of a one-layered idealized arterial segment, reinforced by two fiber families arranged with angles of ±30◦ with respect to the circumferential direction, is exposed to an internal pressure of 16 kpa (120 mmhg). as an approximation of the natural conditions in the c1 [kpa] �1 [kpa] �2 [-] α1 [kpa] α2 [-] 17.5 499.8 2.4 30 001.9 5.1 table 1. material parameters for arterial tissue, adjusted by brands et al. [16] to the model from [9]. ϑ+ [-] k+ϑ [-] m + ϑ [-] ϑ − [-] k−ϑ [-] m − ϑ [-] 1.1 1.0 3.0 0.9 1.0 3.0 table 2. growth parameters used for the simulation of an idealized arterial segment. body, an axial displacement is additionally applied and three different values, namely 10 %, 15 % and 20 % of the initial length, are compared. the material and growth parameters are chosen as given in tables 1 and 2, respectively. the load is applied within 1 s and subsequently held constant until t = 12 s, letting the arterial segment grow. in figure 2, the resulting stress distributions (solid lines) are shown and compared with the stress distributions obtained without including growth (dashed lines). for each value of axial strain, the circumferential stress distribution is significantly smoothened, which reduces the stress peak at the inner surface. the axial stress distributions are smoothened as well and moreover, a distinct stress reduction can be observed which is a consequence of the axial component of growth, compensating the axial deformation. the final axial stress level differs strongly in dependence on the axial displacement: the higher the axial displacement, the higher the axial stresses. examining the growth deformation in terms of the determinant of the growth tensor, depicted in the contour plots in figure 2, it can be observed that the volume change due to growth increases slightly with increasing axial displacement, being nevertheless in the same range. removing the internal pressure while maintaining the axial displacement, the stress distributions given in figure 3 are obtained. being thus not free of external loads in the axial direction, there are non-zero axial stresses in the non-growing reference segments in this state. in the growing segments, non-zero circumferential and axial stresses occur, which are – despite the existence of the axial load – referred to as residual stresses in this contribution. showing circumferential tensile stresses at the outer surface, each of the considered arterial segments is expected to open after a radial cut. the axial stresses in the segments are positive as well, provoking an axial contraction when releasing the axial constraint. the deformations related to these operations are depicted in figure 4. depending on the axial displacement, different opening angles and maximal values of axial strain are obtained. with increasing axial displacement, the opening angle reduces and the axial contraction rises. in the opened segments, the circumferential stresses have vanished. 88 vol. 7/2017 modeling of anisotropic growth and residual stresses in arterial walls -20 0 20 40 60 80 100 120 140 160 180 radial position [mm] 6.5 7 7.5 (a) -20 0 20 40 60 80 100 120 140 160 180 radial position [mm] 6.5 7 7.5 (b) -20 0 20 40 60 80 100 120 140 160 180 radial position [mm] 6.5 7 7.5 (c) 1.34 1.36 1.38 det(fg) 1.33 1.4 radial stress [kpa] circumf. stress [kpa] axial stress [kpa] radial stress, no growth circumf. stress, no growth axial stress, no growth figure 2. radial, circumferential and axial stress distributions as well as growth deformation at t = 12 s in an arterial segment exposed to an internal pressure of 16 kpa and different axial displacements of (a) 10 %, (b) 15 % and (c) 20 % of the initial length. the dashed lines indicate the stresses of a non-growing arterial segment. ra d ia l s tr e s s [ k p a ] -5 -4 -3 -2 -1 0 1 2 3 4 5 radial position [mm] 6.5 7 7.5 (a) c ir c u m f. s tr e s s [ k p a ] -2.5 -2 -1.5 -1 -0.5 0 0.5 1 1.5 2 2.5 radial position [mm] 6.5 7 7.5 (b) a x ia l s tr e s s [ k p a ] -10 -5 0 5 10 15 20 25 30 35 radial position [mm] 6.5 7 7.5 (c) uz = 0.10 l uz = 0.15 l uz = 0.20 l uz = 0.10 l, no growth uz = 0.15 l, no growth uz = 0.20 l, no growth ra d ia l ci rc u m f. ax ia l figure 3. residual stresses in (a) radial and (b) circumferential direction, as well as (c) stresses in axial direction at t = 12 s after removing the internal pressure while keeping up the axial displacement. the stress distributions provoke the axial contraction and opening of a resected, radially sliced arterial segment as it is observed in experiments. (a) (b) (c) -0.5 -0.2 0 0.2 0.5 circumf. stress [kpa] -0.59 0.61 ϕ = 11.06◦ max εz = −1.24 % ϕ = 9.77◦ max εz = −4.63 % ϕ = 8.36◦ max εz = −8.00 % figure 4. deformation states associated to the residual stresses in arterial segments loaded by an internal pressure and axial displacements of (a) 10 %, (b) 15 % and (c) 20 % of the initial length. the wireframe shows the undeformed reference configuration, whereas the second chart form the inside shows the grown state after removal of the pressure. the indicated values of axial strain have been computed with respect to that deformed configuration. 89 anna zahn, daniel balzani acta polytechnica ctu proceedings 4. discussion and conclusion generating favorable results in specific cases as seen in section 3, the performance of the presented model is still strongly dependent on the boundary value problem, especially on the choice of the fiber orientation vectors and on the boundary conditions. the approximation of realistic conditions is difficult for both issues. subjected to remodeling processes, fibers in arterial tissue reorient according to the requirements of their mechanobiological environment. this effect of improving the load-bearing behavior is not included in the current framework so far, such that remodeling caused changes of the growth-inducing stresses and their consequences on growth can not be considered. therefore, the interaction between growth and remodeling processes is a topic which should be taken into account to further improve the model. the axial boundary conditions assumed in this contribution are a vague approach to approximate the natural state of arteries in the body, where axial strains and stresses are existent. thus, one could also have decided to apply axial stresses instead, unknowing which is the best way to mimic the unmeasurable, real conditions. another point with respect to the axial direction shall be mentioned. in contrast to the circumferential direction, where unloading and slicing are two different operations, the compution af axial residual stresses is tricky and demands further meditation. to be able to simulate the contraction after resecting the artery, the applied axial displacement has not been removed in the presented numerical example when computing the “residually” stressed state. doing so, the requested release of the axial constraint could be performed in the modified boundary value problem. if indeed real residual stresses are the matter of interest, the axial loading has to be removed together with the internal pressure, which is only possible when applying an axial stress instead of a displacement. then, the arterial segment can already deform in axial direction during the unloading process, such that no further cutting, revealing the deformation associated to axial residual stresses, is possible afterwards. a third topic for further examinations is the development of a procedure to adjust the growth parameters, which have been chosen arbitrarily until now. acknowledgements the authors appreciate funding from dfg through the institutional strategy “the synergetic university” of tu dresden. references [1] t. matsumoto, k. hayashi. response of arterial wall to hypertension and residual stress. in k. hayashi, a. kamiya, k. ono (eds.), biomechanics. springer, 1996. [2] d. balzani, j. schröder, d. gross. simulation of discontinuous damage incorporating residual stresses in circumferentially overstretched atherosclerotic arteries. acta biomaterialia 2:609–618, 2006. doi:10.1016/j.actbio.2006.06.005. [3] h. weisbecker, d. m. pierce, g. a. holzapfel. a generalized prestressing algorithm for finite element simulations of preloaded geometries with application to the aorta. int j numer method biomed eng 30:857–872, 2014. doi:10.1002/cnm.2632. [4] j. schröder, s. brinkhues. a novel scheme for the approximation of residual stresses in arterial walls. archive of applied mechanics 84:881–898, 2014. doi:10.1007/s00419-014-0838-x. [5] j. schröder, m. von hoegen. an engineering tool to estimate eigenstresses in three-dimensional patientspecific arteries. comput methods appl mech engrg 306:364–381, 2016. doi:10.1016/j.cma.2016.03.020. [6] s. göktepe, o. j. abilez, e. kuhl. a generic approach towards finite growth with examples of athlete’s heart, cardiac dilation, and cardiac wall thickening. journal of the mechanics and physics of solids 58:1661–1680, 2010. doi:10.1016/j.jmps.2010.07.003. [7] p. sáez, e. peña, m. a. martínez, e. kuhl. computational modeling of hypertensive growth in the human carotid artery. computational mechanics 53:1183–1196, 2014. doi:10.1007/s00466-013-0959-z. [8] e. k. rodriguez, a. hoger, a. d. mcculloch. stress-dependent finite growth in soft elastic tissues. journal of biomechanics 27(4):455–467, 1994. doi:10.1016/0021-9290(94)90021-3. [9] d. balzani, p. neff, j. schröder, g. a. holzapfel. a polyconvex framework for soft biological tissues. adjustment to experimental data. int j solids struct 43:6052–6070, 2006. doi:10.1016/j.ijsolstr.2005.07.048. [10] e. kuhl. growing matter: a review of growth in living systems. j mech behav biomed mater 29:529–543, 2014. doi:10.1016/j.jmbbm.2013.10.009. [11] i. hariton, g. debotton, t. c. gasser, g. a. holzapfel. stress-driven collagen fiber remodeling in arterial walls. biomechanics and modeling in mechanobiology 6:163–175, 2007. doi:10.1007/s10237-006-0049-7. [12] s. fausten, d. balzani, j. schröder. an algorithmic scheme for the automated calculation of fiber orientations in arterial walls. computational mechanics 2016. doi:10.1007/s00466-016-1321-z. [13] v. a. lubarda, a. hoger. on the mechanics of solids with a growing mass. int j solids struct 39:4627–4664, 2002. doi:10.1016/s0020-7683(02)00352-9. [14] e. kuhl, r. maas, g. himpel, a. menzel. computational modeling of arterial wall growth. biomechanics and modeling in mechanobiology 6:321–331, 2007. doi:10.1007/s10237-006-0062-x. [15] a. zahn, d. balzani. modeling residual stresses in arterial walls based on anisotropic growth. proceedings in applied mathematics and mechanics 16:115–116, 2016. doi:10.1002/pamm.201610046. [16] d. brands, a. klawonn, o. rheinbach, j. schröder. modelling and convergence in arterial wall simulations using a parallel feti solution strategy. computer methods in biomechanics and biomedical engineering 11:569–583, 2008. doi:10.1080/10255840801949801. 90 http://dx.doi.org/10.1016/j.actbio.2006.06.005 http://dx.doi.org/10.1002/cnm.2632 http://dx.doi.org/10.1007/s00419-014-0838-x http://dx.doi.org/10.1016/j.cma.2016.03.020 http://dx.doi.org/10.1016/j.jmps.2010.07.003 http://dx.doi.org/10.1007/s00466-013-0959-z http://dx.doi.org/10.1016/0021-9290(94)90021-3 http://dx.doi.org/10.1016/j.ijsolstr.2005.07.048 http://dx.doi.org/10.1016/j.jmbbm.2013.10.009 http://dx.doi.org/10.1007/s10237-006-0049-7 http://dx.doi.org/10.1007/s00466-016-1321-z http://dx.doi.org/10.1016/s0020-7683(02)00352-9 http://dx.doi.org/10.1007/s10237-006-0062-x http://dx.doi.org/10.1002/pamm.201610046 http://dx.doi.org/10.1080/10255840801949801 acta polytechnica ctu proceedings 7:85–90, 2017 1 introduction 2 computational framework 2.1 model of anisotropic growth related to two fiber families 2.2 visualization of residual stresses 3 numerical example 4 discussion and conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0063 acta polytechnica ctu proceedings 15:63–68, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app modulus of elascity of unfired rammed earth barbora mužíková∗, tereza plaček otcovská, pavel padevět czech technical university in prague, faculty of civil engineering, thakurova 7, 166 29 prague 6, czech republic ∗ corresponding author: barbora.muzikova@fsv.cvut.cz abstract. the article is focused on measurement and evaluation of elastic modulus of unfired rammed earth and determination of stress-strain curve in compressive test. sets of different containt of clay, sand and water were made and their properties were evalueted and compared. it was found out that for the hightest elascity modulus is ideal the ratio of sand and clay around 75/25 up to 80/20, the sets with montmorillonite clay had substantialy smaller value of elascity modulusthan the sets with illite-kaolinite clay and the ideal content of water for sand/clay ratio 75/25 is around 0.295. keywords: rammed earth, elastic modulus, mechanical properties, stress-strain curve. 1. introduction nowadays utilization of the unfired earth as a building material is rising. the reason of the trend is mainly the fact that unfired earth is according to the principles of sustainable building development. it is environmentally friendly thanks to using of the final product, there is no need to produce any intermediate products. more over unfired earth building is a hundred per-cent recyclable. on the other hand, the mechanical behavior of the material is not described enough to use the data for designing in practical engineering. this was caused by the leaving the rammed earth as a building material in the time when the characteristics of other materials started to be investigated. due to this fact, there are no standards for designing such constructions. that is why the only way how to build such buildings is to have some experience with the material and use gain knowledge in other constructions. nowaday few researches start to pay attention (for example [1–4]) on the topic of unfired earth. worldwide there is only a few companies that are capable to build unfired earth house with all requirements of modern living comfort. one of these constructions is a visitor centre in southwest wyoming (fig. 1). it is built of 2500 m3 rammed earth. the energy is gained from the renewable sources photovoltaic cells, five wind turbines and geothermal system unsuring the heating and cooling of the building. onother example can be a family house in new mexico (fig. 2). there are two big monolithic walls made of rammed earth. over 150 tuns of material were used to build these two walls. the color of layer take turns to look like a rock cliff. there are also very old examples of rammed earth construction (such as the great wall in chine, mayen pyramid in mexico) and even brand new construction are built. there is a map of unfired earth constructions worlwide in the fig. 3. the orange color are areas where the unfired earth is a traditional building material. the white marks are the unfired earth constructions that are stated in the unesco list. as you can see from the fig. 3 unfired rammed earth is used in different climatic zone in the wold [5–10]. figure 1. the tourist centre in wyoming is built of rammed earth [11]. figure 2. the family house in new mexico [12]. 1.1. properties of the material as it was written above rammed earth brings some advantages and of course even disadvantages compared to commonly used building materials such as concrete and steel. 63 http://dx.doi.org/10.14311/app.2018.15.0063 http://ojs.cvut.cz/ojs/index.php/app b. mužíková, t. plaček otcovská, p. padevět acta polytechnica ctu proceedings figure 3. the orange areas are the places where unfired earth is a typical building material and white ones are stated in the unesco list [10]. first of all, lets take a look at the advantages. the material is according to the principles of sustainable development, there are no intermediate products it is direct use of the final product, the building material (in case that there are no stabilisation such as lime or cement) is a hundred procenr recycable and it is a local material. these were the advanteges in the respect of ecology. another properity of unfired earth is that it is able to balances air humidity in rooms. the loam in unfired earth walls is able to absorbe and desorb humidity faster and greater extent than any other materials. from this results keeping very healty living condition with reduced humidity in summer and elevated humidity in winter. the wall of rammed earth (as other heavy materials) stores heat. it can balance indoor climate in areas with hight diurnal temperature differences [13]. another advantages is absorbing pollutants, cleaning the indoor air and shielding high frequency light. we can not forget to mention the fire resistant of unfired earth. the biggest development of massive unfired earth houses building was at the area of the czech republic in the beginning of 17th century. it was caused by big fires that destroyd wooden houses. even so called "fire patent" was published by marie terezie. it was ordering to have brick kitchen and chimney built from unfired bricks [9]. more over earth material preservs timber and other organic materials and it is ideal to do-it-yourself constuction in this case it is inexpensive. on the other hand there are also disadvanteges of the material. from the authors‘ point of view, the main one is a prejudice, skepticism and lack of trustworthy informations about unfired earth worldwive among laymen and even among civil engineers. this is conected to the fact that unfired earth is not a standardises building material. next the material is not water resistant. it must by shetltered against rain and frost, especialy in its wet state. the sheltering can be done by roof overhangs, dampproof courses or appropriate surface coatings [14]. another diadvantage is shringing of the material during its drying [7, 15, 16]. 2. composition of rammed earth the material that we need for creating rammed earth construction is a loam. loam is a product of erosion from rock. the composition and varying properties of loam depend on local conditions. loam is a mixture of clay, silt and sand and sometimes contains larger aggregates like gravel and stones. the particles can be sorted according to diameter: particles with diameters smaller than 0.002 mm are called clay, those between 0.002 and 0.06 mm are called silt, and those between 0.060 and 2.000 mm are called sand. particles of larger diameter are termed gravels and stones. like cement in concrete, clay function is to binde all larger particles in the loam. other larger parts silt, sand and aggregates constitute the fillers in the loam. depending on which of the three components is dominant, we can differentiate clayey, silty or sandy loam. figure 4. the process of building a rammed earth wall. 64 vol. 15/2018 modulus of elascity of unfired rammed earth set sand/clay ratio water-clay ratio pcs of specimens bulk density type of clay [-] [%/%] [-] [-] [kg/m3] [-] gem i 80/20 0.370 4 2073 montmorillionite gem ii 75/25 0.370 6 2040 montmorillionite gem iii 75/25 0.295 6 1929 montmorillionite kr ii 75/25 0.370 11 2134 illite-kaolinite kr iii 80/20 0.400 12 2138 illite-kaolinite s iii 75/25 0.295 4 2170 illite-kaolinite s iv 85/15 0.370 6 2036 illite-kaolinite s v 75/25 0.335 6 2179 illite-kaolinite table 1. composition of used mixtures. 2.1. clay the clay is a product of erosion of feldspar and other minerals. the basic types of clay are known based on the way of creation. these are kaolinite, illite and montmorillionite. they differ in the lamellar structure. the clay changes the properities of the final product mechaical properties and even the colour is different as can be seen in the fig. 5. 2.2. sand sand and gravel are used as fillers. they are described by the graph of grain size distribution. 2.3. water water activates the binding forces of the clay. the amount of mixing water has influence on the properities of the final product. figure 5. different type of used clay in specimens. 3. production of specimens the material for building specimens is loam and mixing water. that means it is a mixture of sand, clay and water. the loam is made in the laboratory, that means the chosen type of clay and sand (of known grain size distribution) are mixed together to create loam of exactly known composition to compare changes in composition with changes of mechanical properties of the final product. the amount of mixing water is determined depending on the amount of clay, and it is expressed by the water-clay ratio. the process of building rammed earth wall is as follows. first of all, the framework is built, then the first layer of rammed earth is filled in. secondly the layer of moist earth is compressed by a tamper. then next layers of moist earth are added and compressed up to the top of framework. finally, the framework is removed and the result is a hard monolithic wall. the principle is shown in fig. 4. eight different prescriptions were tested for determining modulus of elasticity e. sets gem with montmorillonite clay and sets s and kr with illite-kaolinite clay, sets differed in the ratio of sand and clay and in the water-clay ratio. the composition of the prescription is described in table 1. there are names of sets in the first column, then sand/clay ratio, water/clay ratio, the pieces of used specimen is in the forth column, average bulk density of specimens and type of used clay in the last one column. figure 6. testing of specimens: compressive test. the laboratory produce of specimens is similar to the building of real constructions. firstly the prescription is defined as a ratio between sand (the filler) and clay (the binder) and the water-clay ratio is determined. secondly the compounds were mechanically mixed together, then the moist earth en mixture was 65 b. mužíková, t. plaček otcovská, p. padevět acta polytechnica ctu proceedings set s/c w type of clay fc ft e [-] [-] [-] [mpa] [mpa] [mpa] gem i 80/20 0.370 montmorillionite 1.274 ± 0.157 0.072 ± 0.008 43.293 ± 6.865 gem ii 75/25 0.370 montmorillionite 1.211 ± 0.173 0.075 ± 0.007 33.548 ± 7.460 gem iii 75/25 0.295 montmorillionite 1.466 ± 0.433 0.045 ± 0.005 33.528 ± 7.442 kr ii 75/25 0.370 illite-kaolinite 1.150 ± 0.096 54.584 ± 16.856 kr iii 80/20 0.400 illite-kaolinite 1.077 ± 0.104 61.039 ± 22.202 s iii 75/25 0.295 illite-kaolinite 2.017 ± 0.029 0.143 ± 0.009 99.176 ± 12.200 s iv 85/15 0.370 illite-kaolinite 5.890 ± 0.067 0.073 ± 0.005 37.356 ± 17.075 s v 75/25 0.335 illite-kaolinite 2.165 ± 0.037 0.127 ± 0.008 74.339 ± 19.428 table 2. measured modulus of elascity e. layer by layer pressed in to the mould by the drill or manually by a steel block. the specimens were removed from the mould and set in to the climatic chamber where the stable conditions were settled. compressive tests (specimens of size 20×20×100 mm) were made after a month after the produce. the pictures from the testing are shown in the fig. 6 and 7 . figure 7. compresive strenght test with extesometer that records displacements δ on the specimens (range 25 mm). figure 8. principle of determinig of modulus of elascity from σ-ε curve. 4. method of determining e test for determining the static modulus of elasticity was the compresive strenght test. the loaddisplacement curve data were recorded during the test. from these data the stress-strain curve was calculated and then the modulus of elasticity was evalueted. the static modulus of elasticity (so called young‘s modulus) e is a material property, that describes stiffness. mechanical deformation puts energy into a material. the energy is stored elastically or dissipated plastically. the way a material stores the energy is summarized in stress-strain curve (σ-ε curve) that can be determined from load-displacement curve (f-δ curve). stress is defined as force per unit area and strain as elongation per unit length. graphically the modulus of elasticity can be defined as a slope of the linear portion of the stress-strain curve. the modulus of elasticity is defined by the hooke‘s law, this law says that for elastically deformation the stress σ is directly proportional to strain ε: σ = e · ε. (1) the stress-strain curve for gem i set is show in the fig. 8. it can be seen that the elasticity modulus can be determined as: e = ∆σ ÷ ∆ε. (2) the material behaviour is linear elastic, the stressstrain curve in the loading part has a shape of linear curve from the beginning to 0.8fc. due to the fact it was supposed that tangent modulus and secant modulus are of the same value. 5. results from the measurements the arithmetic mean and the standard deviation for each set were settled from the measured values, they are in table 2. there are the compositions of the sets, their compressive and tensile bending stregths ( ± standard deviation) [17–19] and the evalueted 66 vol. 15/2018 modulus of elascity of unfired rammed earth figure 9. prescription of used recepises in triangle diagram. figure 10. comparing values of elascity modulus for all sets. elascity modulus ( ± standard deviation). fig. 10 compares modulus of elascity of each set in a chart where the values are in mpa. the recepis are shown in the triangular diagram (fig. 9) where three components of the mixture (sand, clay and water) are expresed by percent and are on the axis of the diagrame. the composition and the difference between the prescription can bee senn from the diagram. 5.1. impact of sand and clay content the set s iv had the minimum content of clay and water and maximum content of sand (85/15 0.370). let us remember that the amount of water is settle by the water-clay ratio that means the less clay is used the less water is with same water-clay ratio. the clay that was used was the illite-kaollinite clay. this set reached the minimum value of illite-kaolinite clay sets 37.356 ± 17.075 mpa. sets gem ii and kr ii contained the maximum of water and minimum of sand, the ratio is the same at both sets (75/25 0.370). they differed in used clay. illite-kaolinite clay was used at kr ii and montmorillonite at gem ii. the set gem ii and gem iii reached the minimum value of all sets 33.548 ± 7.460 and 33.528 ± 7.442 mpa and the set kr ii reached the second minimum value 54.584 ± 16.856 mpa of sets with illite-kaolinite clay. as a result it seems that the ideal ratio of sand and clay is around 75/25 up to 80/20. 5.2. impact of type of used clay two types of clay were used illite-kaolinite and montmorillonite clay. take a look to already mentioned sets gem ii and kr ii. the ratio of clay, sand and water was the same (75/25 0.370), but the elascity modulus of kr ii with illite-kaolinite clay is 1.6 times higher. another sets gem iii and s iii had the same ratio of clay, sand and water (75/25 0.295) and again they differed in type of clay. illite-kaolinite clay was used at s iii and montmorillonite at gem iii. the elascity modulus of s iii with illite-kaolinite clay is nearly 3 times higher. the sets with montmorillonite clay had substantialy smaller value of elascity modulus so the type of used clay in loam is very important for the mechanical properies of the final product. 5.3. impact of water content the set s iii, s iv and kr ii had the same ratio of sand and clay 75/25. the differed in amount of mixture water 0.295, 0.335 and 0.375. the values of e were 99.176 ± 12.202 mpa, 37.356 ± 17.075 mpa and 54.584 ± 16.856 mpa. so evidently, the optimal value of water-clay ratio is around 0.295 for sand/clay ratio 75/25. 6. conclusions sufficient data of rammed earth properties in relation to its composition have to be find out for extension of these constructions. the article was focused on evalueting elascity modulus of different sets of rammed 67 b. mužíková, t. plaček otcovská, p. padevět acta polytechnica ctu proceedings earth. the main data that were found about dependence elascity modulus and composition of clay, sand and mixing water: • the ideal ratio of sand and clay is around 75/25 up to 80/20, • the sets with montmorillonite clay had substantialy smaller value of elascity modulusthan the sets with illite-kaolinite clay, • the ideal content of water for sand/clay ratio 75/25 is around 0.295. list of symbols e elascity modulus [mpa] f load force [kn] fc comressive strength [kpa] ft tensile bending strength [kpa] δ displacement [mm] ε strain [–] % bulk density [kg m−3] σ stress [kpa] gem sets with montmorillionite clay s sets with illite-kaolinite clay kr sets with illite-kaolinite clay acknowledgements the research was financial supported by the czech science foundation (gacr no.18-0884s) and by the faculty of civil engineering at ctu in prague (sgs project no.16/201ohk1/3t/11). references [1] h. araki, j. koseki, t. sato. tensile strength of compacted rammed earth materials. soils and foundation (56(2)):189–204, 2016. [2] l. miccoli, d. v. oliveira, r. a. silva, u. muller. static behaviour of rammed earth: experimental testing and finite element modelling. materials and structures (48(10)):43–56, 2015. [3] q. liu, t. l. tong. engineering properties of unstabilized rammed earth with different clay contents. journal of wuhan university of technology: materials science (32(4)):914–920, 2017. [4] r. el-nabouch, q. bui, o. ple, p. perrotin. characterizing the shear parameters of rammed earth material by using a full-scale direct shear box. construction and building materials (171):414–420, 2018. [5] b. little, m. morton. building with earth in scotland: innovative design and sustainability. first printing. scottish executive central research unit, 2001. [6] j. norton. building with earth, a handbook. first printing. intermediate technology development group limited, 1986. [7] g. minke. building with earth design and technology of sustainable architecturer. second printing. birkhauser publishe, 2006. [8] p. suske. nepalena hlina v moderni architekture. dokument era21: ekologie, realizace, architektura (2):69–70, 2004. [9] i. zabickova. hlinene stavby. first printing. era vydavatelstvi, 2002. [10] craterre cultures, constructives at developpement durable. 2017 architecture de terre dans le monde. http://craterre.org. [11] sirewall. sirewall inspiration ki,t 2017. http://sirewall.com/. [12] r reck. rammed/earth construction gets luxurz makeover, 2015. http://www.pagethink.com/v/blog-detail/ rammed-earth-construction-is-cool-again-and-hot-/ 47/. [13] x. dong, v. soebarto, m. griffith. strategies for reducing heating and cooling loads of uninsulated rammed earth wall houses. energy and buildings (77):323–331, 2014. [14] q. bui, j. morel, b. v. reddy, w. ghayad. durability of rammed earth walls exposed for 20 years to natural weathering. building and environment (44(5)):912–119, 2009. [15] m. rauch. refined earth construction and design with rammed earth. edition detail. munich, 2015. [16] kianfar, ehsan, v. toufigh. reliability analysis of rammed earth structures. construction and building materials (127):884–895, 2016. [17] t. otcovska, p. padevet. dependence of tensile bending strength of rammed earth on used clay composition and amount of mixture water. modern methods of experimental and computational investigations in area of construction ii pp. 48–53, 2017. [18] t. otcovska, p. padevet. dependence of compressive strength of rammed earth on used clay composition. experimental stresss analysis 2016. [19] i. zabickova, t. otcovska, p. padevet. compressive strength of unburned clay masonry. modern methods of experimental and computational investigations in area of constructions pp. 31–34, 2016. 68 http://craterre.org http://sirewall.com/ http://www.pagethink.com/v/blog-detail/rammed-earth-construction-is-cool-again-and-hot-/47/ http://www.pagethink.com/v/blog-detail/rammed-earth-construction-is-cool-again-and-hot-/47/ http://www.pagethink.com/v/blog-detail/rammed-earth-construction-is-cool-again-and-hot-/47/ acta polytechnica ctu proceedings 15:63–68, 2018 1 introduction 1.1 properties of the material 2 composition of rammed earth 2.1 clay 2.2 sand 2.3 water 3 production of specimens 4 method of determining e 5 results from the measurements 5.1 impact of sand and clay content 5.2 impact of type of used clay 5.3 impact of water content 6 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.10.0052 acta polytechnica ctu proceedings 10:52–55, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app testing of soil improvement below a hall floor monika súľovská∗, peter turček, zuzana štefunková stu bratislava, faculty of civil engineering, radlinského 11, bratislava, slovakia ∗ corresponding author: monika.sulovska@stuba.sk abstract. construction of a hall situated in north-west slovakia was prepared in winter. after finishing an elementary ground adaptation in february there were extreme weather conditions: the precipitation exceeded by over 300 % the normal long term measured amount. this indisposition of climate led to a stopping of the construction work. checks on the required quality of the base course using static load tests showed a failure in performance of compaction. the sampling and laboratory testing of the original and the improved soils are discussed in the paper. keywords: quality control, base course, laboratory analysis, consolidation test, cement content. 1. introduction the required intensity of deadlines for finishing the construction regardless of limiting conditions (e.g. climatic, occurrence of soils with low bearing capacity) conduced on a good number of occasions to improvement efforts on the base course. these are usually similar technologies to those used in road construction. among the most common technologies there is included adding a certain percentage of lime or cement to fine-grained soils6 (so-called soil stabilization). mixing cement with soil and compaction reduces the compressibility of the treated soil. subsequent quality control of any improvement is mostly done by static load test. the underlying layers of the future floors were simply improved within the rainy weather and were not protected. the adverse weather situation caused the suspension of work. checking the quality of the underlying layers using static load tests after a week of completion of the stabilization showed a failure to satisfy the required improvement. it was necessary to immediately resolve the situation and help meet deadlines. consequently soil was taken from the unsatisfactory area of the hall and then a check was made on the properties of both treated and untreated layers by laboratory tests. 2. laboratory tests on the construction site of the hall there were removed undisturbed soil samples improved by cement, but also the original soil without improvement by cement. on the original soil samples there were firstly detected the basic descriptive soil characteristics. based on stn 72 1001 [1] there was soil classified as firm clay with low plasticity (f6 cl). the original moisture of the soil was more than 20 %. the moisture of the improved soil with cement did not differ from the moisture of the original soil. to determine the optimum moisture of soil for compaction there was used the proctor test (see fig. 1). the optimum moisture of the tested soil was determined wopt = 15.2 %. this corresponds to the maximum dry density of soil ρmax = 1.768 g/cm3. the collected samples had significantly higher moisture content than its optimal moisture. 2.1. oedometric tests on undisturbed soil samples compressibility testing on the oedometer was made to the original and also the improved soil. at improvement there was added into the soil 5 % of cement. the results of average deformation characteristics of original and improved soil are collected in table 1. on the basis of laboratory results there can be determined for the improved soil the following increases in deformation characteristics: • at stresses in soil σ ≤ 50 kpa eoed stab eoed not = 32.25 5.10 = 6.32 (1) • at stresses in soil 50 ≤ σ ≤ 100 kpa eoed stab eoed not = 14.13 5.70 = 2.47 (2) • at stresses in soil 100 ≤ σ ≤ 200 kpa eoed stab eoed not = 15.19 9.62 = 1.58 (3) based on the stress range of its own weight and the expected load from buildings, the improvement of soil after adding cement was 1.5 to 2 times. the recommended values in the deformation modulus of improved soil should not exceed the value eoed = 14 mpa. 2.2. compression tests of soil without improvement at compaction after the proctor standard artificially prepared samples of soil compacted after proctor standard were tested in the oedometer. the 52 http://dx.doi.org/10.14311/app.2017.10.0052 http://ojs.cvut.cz/ojs/index.php/app vol. 10/2017 testing of soil improvement below a hall floor figure 1. determination of the maximum bulk density versus optimum moisture. table 1. deformation modulus of soils determined from compression tests. resulting average oedometer modulus corresponding to different moistures of tested soils are evaluated in table 2. effects of changes in soil moisture on the oedometric modulus are shown in fig. 2. the plasticity index of soil was ip = 10 % (liquid limit wl = 30.5 % and plasticity limit wp = 20.5 %). until the soil was in natural conditions and firm consistency, there were found very small values of deformation characteristics of tested soil to all extents of normal stresses. increasing the soil stiffness (higher values of eoed) was possible by observing a decreasing of soil moisture below 17 %. the consistency of soil at such moisture was ic = 1.4. decreasing the moisture below 17 % significantly increases the oedometric modulus of the soil. analysis of fig. 2 taking into account fig. 1 can be stated as: • exceeding the optimum moisture wopt occurred in samples loaded by normal stress of 50 kpa with rapid decreasing of eoed. • with an increasing of moisture there was observed a significant decreasing of eoed at lower normal stress (< 50 kpa) more than 5-times. exceeding the value of normal stress 50 kpa and with increasing the moisture there was found a decreasing of eoed approximately 3.5 times. in the test range of normal stress eoed was significantly affected by moisture. in fig. 3 there is shown the relationship between deformation characteristics and normal stresses of soil at various initial moistures of the soil. also in this evaluation it was proved the moisture was around 17 % than the limit. at lower moisture there were detected significantly different deformation parameters, especially at normal stress to 50 kpa. 2.3. improving the soil characteristics by adding cement at natural moisture of soil 20 % of collected samples had values of deformation modulus of the original soil which were very low. the addition of cement decreased soil moisture, whilst increasing the deformation parameters was not sufficient. very interesting results were obtained from evaluation of the effect of the 5 % addition of cement (see fig. 4). at lower humidity than 20.5 % (on a stiff consistency) there was observed a significant effect of the cement. at the same time, the effect of improvement increases with lower levels of normal stress. adding 5 % of cement at optimal soil moisture brought a 2-3 times increase of eoed compared to unimproved soil. exceeding soil moisture w = 20.5 % the soil fell into firm consistency and the impact effect of the cement was shown to be very low to negligible. cement added to the soil with natural moisture, got between 19 to 21 %, extracting a portion of water by hydration. this will inevitably reduce soil moisture itself. it can be assumed that the excess of the initial 53 m. súľovská, p. turček, z. štefunková acta polytechnica ctu proceedings table 2. average oedometric modulus of tested soils at various moistures. figure 2. relationship between moisture of tested soil and oedometric modulus. moisture content in excess of the wopt + 5 % there is to be expected a significant decrease in the effect of stabilization. 3. the quality assessment and the cement content in the soil cement and its contents were independently tested in the stabilized soil. laboratory tests of the physicalmechanical properties of cement were carried out and supplemented by x-ray and dta analysis. stabilized soil was investigated by dta and also x-ray analysis. tested cement was used to stabilize the soil at the construction site. the studied properties and the results derived are as follows: • after 28 days, the compressive strength reached 48.3 mpa. this value is in accordance with the requirements of the class of cement 42.5 mpa according to en 197-1. • bulk density with a timed deposit of test samples slightly increased. after 28 days, the bulk density reached a value around 2190 kg/m3, which corresponded with standard values. • by testing the cement slurry, it was found that the cement needs to reach a slurry of normal density of 32 % water. • the initial set of cement slurry observed after 195 minutes, in which it fulfilled the limit set by the standard. final set occurred after 250 minutes. • soundness of the testing cement went to 6.0 mm, which met the standard set limit of 10 mm. • bulk density of fresh cement mortar was 2160 kg/m3. all testing samples had low scatter from the average. • x-ray analysis showed the incidence of cement clinker minerals as alite c3s (tricalcium silicate), belite c2s (dicalcium silicate), brownmillerite c4af (tetracalciumaluminoferrite), c3a (tricalcium aluminate) and gypsum setting regulator (caso4 · 2h2o). furthermore, there occurred silica (sio2), calcite (caco3) and portlandite ca(oh)2. according to this composition it can be assumed that this was portlandcomposite cement cem ii. tests of the stabilized soil by method x-ray and dta showed that samples taken from the hall after the defective area of the static load tests contained less than 2 % of cement. by applying the same method to artificially prepared samples there was found good correspondence with the percentage of added cement. 4. conclusion based on the compressibility tests on the stabilized soil it is clear that the high moisture of the soil after the rainy season had not reached sufficiently an increasing of the soil parameters. they showed two reasons for failure to satisfy the criteria required for static load testing. the first reason was lower cement content as determined by the project. instead of 5 % of the cement which was in samples taken from the 54 vol. 10/2017 testing of soil improvement below a hall floor figure 3. dependence of normal stress and moisture on deformation characteristics of soil. figure 4. influence of addition 5 % cement into the soil on eoed at optimum moisture. subgrade in the construction site there was found only a cement content in the range of 1 to 2 %. significant precipitation caused the natural soil moisture to exceed 20 %. laboratory experiments have shown that at such a level of moisture in the not improved original soil the deformation properties of soil are very low. mixed cement decreased soil moisture, but it reflected only a small increase of eoed. deformation parameters of soil by adding cement may bring an increase 2-3 times. this effect is linked to the initial soil moisture before stabilization; value wopt should not be exceeded by 5 %. in exceeding this moisture there is not expected a major stabilization effect. 5. acknowledgement the paper is a part of results from a research project supported by the slovak agency vega no. 1/0882/16 “influence of boundary conditions to limit states of geotechnical structures”. [2] [3] [4] [5] [1] references [1] stn 72 1001: classification of soil and rock in engineering geology and geotechnics (in slovak), 2010. [2] p. turček, m. súľovská, z. štefunková. evaluation of soil layers improving below the production hall (in slovak). svf stu bratislava, 2016. [3] p. turček, m. súľovská, z. štefunková. experimental research of deformation properties of fine-grained soil situated in hall subgrade (in slovak). svf stu bratislava, 2016. [4] p. turček, j. frankovská, súľovská. geotechnical design according to eurocodes (in slovak). engineering advisory centre of slovak chamber of civil engineers, 2010. [5] en 197-1: cement. part 1: composition, specifications, and conformity criteria for common cements, 2012. 55 acta polytechnica ctu proceedings 10:53–56, 2017 1 introduction 2 laboratory tests 2.1 oedometric tests on undisturbed soil samples 2.2 compression tests of soil without improvement at compaction after the proctor standard 2.3 improving the soil characteristics by adding cement 3 the quality assessment and the cement content in the soil 4 conclusion 5 acknowledgement references acta polytechnica ctu proceedings doi:10.14311/app.2018.17.0020 acta polytechnica ctu proceedings 17:20–23, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app fitexc – diffraction profile fitting program run in ms excel petr vertata, b, ∗, jan drahokoupila, c a department of solid state engineering, fnspe ctu in prague, trojanova 13, 120 00, prague 2, czech republic b department of magnetic measurements and materials, institute of physics as cr, na slovance 2, 182 21, prague 8, czech republic c department of material analysis, institute of physics as cr, na slovance 2, 182 21, prague 8, czech republic ∗ corresponding author: vertat@fzu.cz abstract. we present our custom made diffraction fitting program fitexc for the basic treatment of the 1d diffraction profiles, in particular, the intensity(2θ) of x-ray, synchrotron or neutron diffraction measurements. in the current state, the program allows for simultaneous fitting of up to three profiles in one data set and is intended to be used mainly for precise lattice parameters determination or fitting of overlapped peaks. the entire program is built in the common ms excel environment, it does not require any installation and is free to use. keywords: x–ray diffraction, diffraction profile fitting, ms excel. 1. introduction x-ray crystallography is a powerful tool with many applications: it is often used to determine the lattice parameters, the complete structure of solids, to examine the phase composition, residual stresses or texture. for multiple applications the precise examination of measured 1d intensity(2θ) diffraction profiles represents a crucial part of data processing. there are various programs for detailed diffraction profile fitting available. many of them offer complex solutions implementing the rietveld refinement process [1] (such as x’pert highscore plus [2], topas [3] or fullprof [4]). on the other hand, there are many applications, where quick and precise fitting of few diffraction peaks and their closest neighbourhood is sufficient. even in this field, a few programs are used, such as fityk [5]. unfortunately, the majority of the programs for x-ray diffraction (xrd) data analysis does not, to our knowledge and experience, allow the user to fully control or customize all the aspects of the fitting in an easily accessible way, nor are they userfriendly or quick enough for easier applications and batch processing. although the simple fitting of a few diffraction maxima might also be done manually (for example using the origin software [6]), the approach is often clumsy and require wider knowledge of such environment. we developed a simple user-friendly diffraction profile fitting program fitexc that does not suffer from mentioned disadvantages, yet provides precise results for basic applications. the program, first introduced in [7], is based on the microsoft excel environment, which is immediately available on the majority of computers today. excel’s interface together with visual basic for applications-based macros create a lightweight but powerful tool for the diffraction peak fitting. another advantage is the immediate possibility to implement user’s own scripts for various kinds of analyses such as macroscopic residual stress determination or the analysis of the crystallite size and strain using the williamson-hall plot. 2. program interface and fitting functions in the attempt to make fitexc as user-friendly as possible, the clickable interface with buttons, drop-down lists and check-boxes was created over the hidden worksheets used for the calculations. actions are macro-powered and the overall approach is straightforward. first, user is asked to select the file to be loaded. panalytical .xrdml files are supported as well as standard .xy, .xye or .txt files (multiple columns in the ordinary textfile). next, the fitting function and the radiation spectrum can be specified. standard cauchy, gauss, pearson vii and pseudo-voigt functions are implemented, as well as their split variants for dealing with possible asymmetry. as for the radiation, user may choose from different x-ray tube presets or define his own spectrum containing up to 7 components with desired wavelength, relative intensity and width (figure 1) [8]. therefore, the program might be used not only for xrd, but also for neutron or synchrotron diffraction data fitting. on the main fitting page, where the measured data and fitted function are plotted (figure 2), the entire data processing is controlled. in the current state, up to three profiles might be fitted for one loaded dataset, in the future, the ability to use more than three profiles is to be added. each of the profiles is defined by 20 http://dx.doi.org/10.14311/app.2018.17.0020 http://ojs.cvut.cz/ojs/index.php/app vol. 17/2018 fitexc – diffraction profile fitting program run in ms excel figure 1. program interface: initial settings worksheet shown after the data file is chosen. here the basic settings can be specified, such as the spectrum of the radiation, function for fitting or initialisation of the new peak parameters. function formula (f(2θ)) cauchy a1/(1 + a3(2θ − a2)2) gauss a1exp(−a3(2θ − a2)2) pearson vii a1/(1 + a3(2θ − a2)2)a4 pseudo-voigt a4 gauss + (1 − a4) cauchy table 1. implemented fitting functions with following parameters: intensity at maximum a1, angular position of the maximum a2 , the relative width a3 and the shape parameter a4 [9]. four parameters a1−4, where a1 is the intensity at maximum, a2 the 2θ position of the maximum, a3 is the relative width parameter and a4 is the shape parameter. the simple formulas for such defined functions are denoted in table 1 [9]. in the case of asymmetric profiles, instead of one a3 and a4 parameter, two sets of parameters (a31, a41 and a32, a42) are used – one pair for the left and one pair for the right side of the peak. for background fitting, the constant, linear and quadratic forms were implemented so far. the program also calculates the integral intensity of the fitted reflections using analytical formulas where possible and numerical approach for the pearson vii based functions. the loaded 2θ might be restricted if needed. as one of the advantages, the user is immediately able to fix some parameters that will not be changed during fitting. this makes the program very effective for fitting of complex patterns and multiple overlapped peaks. fitting algorithm is based on general reduced gradient non-linear least squares method. it is implemented using the solver (solver.xlam) add-in and the standard worksheet environment of ms excel run in the background. to evaluate the goodness of the fit the standard discrepancy factors rp , rw p , rexp and χ2 are calculated and their values might be compared with those before fit. 3. additional functionality for processing of multiple data, our program allows loading of multiple files. these are then fitted automatically one after another. in case that .xrdml file containing multiple scans is loaded, the batch feature is enabled automatically providing the tool to fit the datasets sequentially. this feature has already been used for determination of the temperature development of selected lattice parameters of magnetic shape memory (msm) alloys [10]. as our program enables determination of very small changes of the lattice parameters, their temperature evolution is clearly observable even in narrow temperature intervals (figure 3). fitting algorithm is stable enough even for fitting of multiple peaks with great intensity differences 21 petr vertat, jan drahokoupil acta polytechnica ctu proceedings figure 2. program interface: the main worksheet for data fitting. user is able to start the fitting, preset or check the fitted parameters, fix them if needed and to see the measured and fitted data including the difference curve. figure 3. application example: precisely determined lattice parameters a and b of ni–mn–ga msm alloy and their evolution while heating. the sudden change denotes the structural transformation [10]. (so far even with two orders difference the results were reliable). to review a situation, the user may immediately choose between the linear and logarithmic scale of the main graph by clicking a single button. once the good fit is obtained, all the results including the basic settings might be stored in the easily accessible results worksheet. the procedure for easy recovery of the parameters of the fit to the fitting worksheet is also implemented. 4. conclusions we developed a new simple diffraction profile fitting program fitexc. this lightweight and user-friendly program can be run on any pc with ms excel (2003 or later) installed. it is intended mainly for the precise fitting of the single reflections (especially their 2θ positions) based on selected or user-defined spectra of the radiation source and for fitting of overlapped peaks. this allows the precise determination of the lattice parameters, observation of their evolution or study of the residual stresses. the program is free to use. to get your own copy, please visit http://people.fjfi.cvut.cz/vertapet/. acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs16/245/ohk4/3t/14, and czech science foundation, grants no. 16-00043s and 17-00062s. authors would like to thank o. heczko for useful remarks and motivation the tremendous amount of tasks requiring the development of our program :-) . references [1] h.m. rietveld: a profile refinement method for nuclear and magnetic structures. j. appl. cryst. 2 65–71, 1969. doi: 10.1107/s0021889869006558 [2] t. degen, m. sadki, e. bron, u. koenig, g. nenert: the highscore suite. powder diffraction 29 (supplement s2):s13–s18, 2014. doi: 10.1017/s0885715614000840 [3] diffrac.suite topas (bruker, karlsruhe, germany), https://www.bruker.com/products/x-raydiffraction-and-elemental-analysis/x-raydiffraction/xrd-software/topas.html 22 vol. 17/2018 fitexc – diffraction profile fitting program run in ms excel [4] j. rodriguez–carjaval: recent advances in magnetic structure determination by neutron powder diffraction. physica b 192 (1–2):55–69, 1993. doi: 10.1016/0921-4526(93)90108-i [5] m. wojdyr: fityk: a general-purpose peak fitting program. j. appl. cryst. 43 1126–1128, 2010. doi: 10.1107/s0021889810030499 [6] origin (originlab, northampton, ma). [7] p. vertat: diffraction profile fitting in ms excel. 5. studentska vedecka konference fyziky pevnych latek, nove hrady, ceske vysoke uceni technicke v praze, 2015. [8] v. k. pecharsky, p .y. zavalij: fundamentals of powder diffraction and structural characterization of materials, kluwer academic publishers: boston/dordrecht/london , 2nd ed., p. 115, 2009. [9] r. kuzel: informace v praskovem difraktogramu a jeho zpracovani. materials structure 10 (1a) 18–19, 2003. [10] l. straka, j. drahokoupil, p. vertat, j. kopecek, m. zeleny, h. seiner, o. heczko: orthorhombic intermediate phase originating from {110} nanotwinning in ni 50.0 mn 28.7 ga 21.3 modulated martensite. acta mater. 132 335–344, 2017. doi: 10.1016/j.actamat.2017.04.048 23 acta polytechnica ctu proceedings 17:20–23, 2018 1 introduction 2 program interface and fitting functions 3 additional functionality 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0038 acta polytechnica ctu proceedings 5:38–40, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app selected aspects of etcs-l3 deployment petr koutecký czech technical university in prague, faculty of transportation sciences, department of transport telematics, konviktská 20, prague, czech republic correspondence: koutepe1@fd.cvut.cz abstract. this paper focuses on selected problem areas in the implementation of etcs l3, analyz-es their impacts and offers a number of possible solutions as a basis for pilot testing. it compares the suitability of each solution and summarizes their advantages and disadvantages. keywords: ertms, etcs l3, odometry, its-r. 1. introduction control and command systems on the railways have undergone a relatively long development. on the one hand, the development is driven by increasing demands on safety, reliability, smoothness and fuel economy, on the other hand, the development was mainly limited by technical capabilities of the time in the past. the development possibilities of electronics and especially microprocessor technology and computing power have allowed in recent years to increasingly address the question of optimizing the control process quality and minimizing the necessity of the human factor’s intervention, which in earlier times was not possible to such an extent. this technical progress, pressure on cost cutting and the european union requirements for the interoperability of the railway networks have created the need for addressing the management and security of railway services in new ways. introducing new systems brings the need of identifying and addressing the potential risks and technical challenges associated with the transition to the new system. the aim of this paper is to point out possible problems in the implementation and deployment of etcs l3 on the conventional railway network and suggest possible solutions to these problems with the specification of advantages and disadvantages of the various options. 2. requirements for the full implementation of intelligent transport systems on railways (hereinafter its-r) and for the purposes of this article, we consider the target state to be a state in which the following conditions are met: (1.) there is a bidirectional data connection to the train (between mobile and stationary parts), (2.) the integrity check of the train using technical means on the vehicle is ensured, (3.) all locomotives are equipped with an on-board unit (obu) etcs. when projecting these demands while considering the current trend and the commitments of the european union member states in the field of the railway network interoperability, the ertms system on application level 3 (l3) is a suitable candidate for addressing the command and control system (ccs). 2.1. substantiation of the requirements requirements specified in the previous section are based primarily on the overall concept of its-r. the need for a data connection to the train is given mainly by the necessity of the existence of such a data transfer connection between the stationary and mobile parts of the etcs l3 system. however, the data connection can also be used with other subsystems in the its-r concept, such as the transmission of data for the ato system, transmission of diagnostic data, etc. another requirement – to ensure the train integrity – is another necessary condition for the etcs l3 system’s activity, where the existence of detection devices already disregarded. the third requirement requiring full equipment of the cabin section in all vehicles is again given by application level 3 of the etcs system, where mixed traffic (such as with l2) is operationally unacceptable after removing the signals. only the connection and fulfillment of all three requirements above at the same time gives us the opportunity to fully start taking advantage of etcs l3 âăş primarily to use the opportunity to minimize the infrastructure elements (signals, track circuits, axle counters). such minimization has several far-reaching impacts – reducing the investment demand compared to the situation where application level 2 (or even one) of the etcs system is installed to achieve the desired safety parameters; according to [1] by approximately 25% compared with l2. it is necessary to emphasize that etcs l1 and l2 are "only" an extension of conventional interlocking equipment and block signaling system and track security equipment. in practice this leads to duplication of some functions in rbc and ixl. this delivers increased investment and operating costs for the construction and operation of two parallel systems and also reduces the possibilities of 38 http://dx.doi.org/10.14311/app.2016.5.0038 http://ojs.cvut.cz/ojs/index.php/app vol. 5/2016 selected aspects of etcs-l3 deployment figure 1. required odometry accuracy [2] the system – especially if a link rbc -> ixl is not implemented for l2. only etcs l3 is not an extension of the current ixl – it is thus possible to leave out the riding principle in a fixed block and train ride control using the signals. 3. definition of problem areas 3.1. the accuracy of the odometer a key part of the function of determining train location in the etcs system is an odometer and doppler radar. since this is a safety-relevant function, it is necessary to consider a possible error in the given position in a safer direction. according to the specification [2], the required accuracy is expressed by the formula according to the traveled distance s: sdif f = 5 + 0.05xs[m] the position is specified (error is null) always after reading a balise group and the error refers to the distance traveled from that point. for the operation of the etcs system in its current form it is therefore necessary to count with this error in determining the usable length of each track. in addition, the position of the end of the train must be added to this inaccuracy in the case of etcs l3, e.g. for track length of 600 m it is necessary to consider an error of up to 35 m, and vacate the space behind the railroad switch signal at least by this distance in order for the railroad switch stretch to be regarded as vacant. this again reduces the usable length of the track. 3.2. train handling the method for determining train location and sections occupancy in the stretch of etcs l3 makes it necessary to address potential problems arising from the train handling, such as splitting, joining and shunting of train cars. the simplest situation occurs when joining and separating sets of coherent units, where with regard to the assumed equipment of all driving vehicles it will be possible to locate both parts and detect their integrity after division. however, this does not apply to the movement of parts of the train set and keeping a part of the train on the station track or to the train profiling into multiple parts. in such a situation, the vacancy of the concerned part of the infrastructure cannot be safely determined only with the use of the etcs l3 system’s technical means. 4. suggested solution 4.1. the accuracy of the odometer to increase the usable length of the transport rails, it may be necessary to increase the accuracy of the positioning system. with current specifications, this problem can be solved by placing one or more locators inside the balises of a long track section. this will split it into several smaller sections, where the odometer’s inaccuracy is proportionately smaller. another possibility is the gradual tightening of the requirement for maximum permissible error in the etcs odometer (now 5m + 5%). similar devices used in the cbtc systems achieved accuracy of about 1% according to [3] already (using e.g. a doppler radar with multiple antennas). 4.2. train handling the problem with train division, shutting down of train cars, and shifts can be solved in principle several ways, which will be discussed in the following text. the advantages and disadvantages of various options and a comment on the possible applicability are also mentioned. 4.2.1. confirmation of occupancy by human operators the occupancy of selected parts of the infrastructure can be confirmed with the participation of the human factor. this principle is used according to [4], e.g. in ertms regional in sweden. after confirming the section’s occupancy with the human factor, the ride through the stretch is possible only in the staff responsible mode with speed restrictions. after the first train passes, the stretch is already considered vacant from the point of view of previous manipulation, and the ride of the following trains is no longer restricted. the disadvantages of this solution are obvious – the need for participation of the human factor, the risk of failure is limited by limiting the speed for the next train. this thereby decreases the throughput performance parameter. however, this solution may be suitable for regional routes with limited speed and sporadic shifts. 4.2.2. partial equipment with detection resources in case of frequent handling, using conventional detection devices (track circuits, axle counters) at this point is offered as a solution to the question of detecting the occupancy of the infrastructure’s section. the disadvantage regarding the involvement of the human factor on detecting the occupancy of the section is thus eliminated, the disadvantage being the need to install detection devices, albeit to a very limited extent. 39 petr koutecký acta polytechnica ctu proceedings 5. evaluation, conclusion the article pointed out several questions that will be necessary to address within the preparation of the possible pilot deployment of etcs l3. the solution’s suitability in some cases depends on the local situation and the nature of the track on which the system will be deployed; all possible situations cannot generally always be covered. in the near future it is also possible to expect verification of the related technologies that could help contribute as another alternative to solving open points (e.g. use of satellite positioning systems instead of fixed balises, use of modern communication technology instead of gsm-r, etc.) in the event of a positive outcome. 6. abbreviation • ato ... automatic train operation • cbtc ... communication based train control • ccs ... command and control system • ertms ... european rail traffic management system • etcs ... european train control system • gsm-r ... global system for mobile communications-railway • its-r ... intelligent transport system-railway • ixl ... interlocking • obu ... on-board unit • rbc ... radio block centre references [1] w. coenraad. level 3 : from high speed vision to rural implementation [online], 2011. http://www.irse.org/knowledge/publicdocuments/ 2011_09_29_ertms_level_3_from_high_speed_ vision_to_rural_implementation.pdf. [2] ertms/etcs performance requirements for interoperability, subset-041, issue 3.1.0, 2012, unisig [online]. http://www.era.europa.eu/document-register/ documents/set-2-index014-subset-041%20v310.pdf. [3] ertms level 3 risks and benefits to uk railways, final report; transport research la-borator [online], 2010. http: //www.trl.co.uk/media/773946/cpr798_-_ertms_ level_3_risks_and_benefits_to_uk_railways.pdf. [4] ertms regional functional requirements specification, version 03.08, uic, 2009. 40 http://www.irse.org/knowledge/publicdocuments/ 2011_09_29_ertms_level_3_from_high_speed_vision_to_rural_implementation.pdf http://www.irse.org/knowledge/publicdocuments/ 2011_09_29_ertms_level_3_from_high_speed_vision_to_rural_implementation.pdf http://www.irse.org/knowledge/publicdocuments/ 2011_09_29_ertms_level_3_from_high_speed_vision_to_rural_implementation.pdf http://www.era.europa.eu/document-register/documents/set-2-index014-subset-041%20v310.pdf http://www.era.europa.eu/document-register/documents/set-2-index014-subset-041%20v310.pdf http://www.trl.co.uk/media/773946/cpr798_-_ertms_level_3_risks_and_benefits_to_uk_railways.pdf http://www.trl.co.uk/media/773946/cpr798_-_ertms_level_3_risks_and_benefits_to_uk_railways.pdf http://www.trl.co.uk/media/773946/cpr798_-_ertms_level_3_risks_and_benefits_to_uk_railways.pdf acta polytechnica ctu proceedings 5:38–40, 2016 1 introduction 2 requirements 2.1 substantiation of the requirements 3 definition of problem areas 3.1 the accuracy of the odometer 3.2 train handling 4 suggested solution 4.1 the accuracy of the odometer 4.2 train handling 4.2.1 confirmation of occupancy by human operators 4.2.2 partial equipment with detection resources 5 evaluation, conclusion 6 abbreviation references acta polytechnica ctu proceedings doi:10.14311/app.2018.20.0073 acta polytechnica ctu proceedings 20:73–77, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app regarding certain aerodynamic sources of non-stationary force loads on turbine profiles tomáš radnic∗, martin luxa, david šimurda institute of thermomechanics of the czech academy of sciences, dolejškova 1402/5, prague, czech republic ∗ corresponding author: radnic@it.cas.cz abstract. the paper deals with selected phenomena present in a blade cascade flow field. the presented research was performed on prismatic blade cascades composed mostly of the tip and root sections of the last stage rotor blade of a large output steam turbine. the unstarted supersonic flow on the cascade inlet, the separation of the flow and the swirl line behind the trailing edge have been identified as the possible sources of the unsteady force effects. keywords: turbine blade cascade, unstarted supersonic flow, swirl line, off design conditions. 1. introduction the modern steam turbines of large power output utilize very long blades in the last stage of the rotor. the tip of the blade operates at supersonic inlet velocities, while the root of the blade operates at subsonic inlet velocities. in this setup, it is certain that the transonic inlet flow conditions will occur on the blade at its nominal speeds, hence unstarted supersonic flow will always be present. during the start up or shut down of the turbine, off design conditions appear in the turbine, which can generate swirl lines and flow separation. although these phenomena can only be present during off design conditions, they can prove fatal for the turbine. last but not least, in modern times, renewable sources are widely incorporated into the energetic grids and their non-reliable power output requires turbines to be able to operate in off design conditions on standard basis. it is important to note the majority of phenomena leading to unsteady force effects occurs in the region of tip section which is not rigid. 2. experiments the investigated phenomena were measured on different sections of a long turbine rotor blades made by doosan skoda power, formerly škoda pilsen. the investigation of the flow field is usually performed on several sections of the blade. tip sections and root sections were used for creation of prismatic models that could be fitted into test section of the wind tunnel. the prismatic blades cascades with finite number of blades were assembled for maintaining best possible periodicity of the flow field that would models as best as possible the flow field in the real turbine. perforated tailboard was utilized during measurements to prevent reflection of the shock waves and help maintaining periodicity. blades with tailored trailing edge with sharp edges, see figure 1, were used for investigation of the swirl line behind the trailing edge. figure 1. one of the shapes of the trailing edge used for the swirl line investigation. figure 2. scheme of the supersonic indraft wind tunnel. 1) silica gel drier 2) pebble filter 3) inlet nozzle 4) adjustable supersonic nozzle 5) test section 6) settling chamber 7) control nozzle 8) quick acting valve 9,10) duct to vacuum chamber. measurements were conducted in the supersonic indraft wind tunnel of the institute of thermomechanics of the czech academy of sciences in nový knín. the tunnel is equipped with deformable inlet nozzle so that supersonic inlet speeds could be achieved. 3. unstarted supersonic flow the unstarted supersonic flow can be characterized as a regime, where the inner branch of inlet shock wave has a shape of a normal shock, see figure 3. when the inner branch of the shock wave turns into chevron shape, the flow is considered to be started supersonic flow, see figure 4. the unstarted supersonic flow is also a boundary between transonic and supersonic flow past the cascade. in case of the unstarted supersonic flow, the inner branch of the inlet shock wave 73 http://dx.doi.org/10.14311/app.2018.20.0073 http://ojs.cvut.cz/ojs/index.php/app t. radnic, m. luxa, d. šimurda acta polytechnica ctu proceedings figure 3. unstarted supersonic flow field. figure 4. started supersonic flow field. has always unstable and unsteady behaviour and can induce instability of the whole flow field [1]. the transition of unstarted flow to started flow can be achieved by increase of inlet speed while maintaining low back pressure or by reduction of backpressure if the inlet speed is high enough. it is important to note that the unstarted supersonic flow cannot be avoided in the turbine where the tip of the blade operates at supersonic inlet velocities. the phenomena will be present on the blade and will have force effect on the blade. the force effects need to be assessed to reckon with them during the design phase of the turbine. the transition of unstarted supersonic flow to started supersonic flow can be seen in figure 5. the outlet isentropic mach number was roughly (a) . m1 = 1.141 m1 = 1.180. (b) . m1 = 1.209 m1 = 1.209. figure 5. development of the started supersonic flow. figure 6. scheme of torque load of the tip of the blade with non started supersonic flow field. m2 is ≈ 1.8. in the top left image, the inlet mach number is m1 = 1.141. the inlet shock wave has a shape of a normal shock wave that interacts with the boundary layer on the pressure side of the adjacent blade. the interaction causes local separation of the flow. on the top right image there is m1 = 1.180. the inlet shock wave is now much closer to the leading edge and the local separation which it causes is now much more limited. on the bottom there are two images with identical inlet speed m1 = 1.209. the images show unstable nature of transition between unstarted and started supersonic flow. on the left image the inner branch of the inlet shock wave hits the trailing edge of the adjacent blade, while on the right image the inlet shock wave has chevron shape. the inner branch of the inlet shock wave on the right image does not collide with the adjacent blade at all, however it can still affect the area of near wake. the shift of the shock wave shape does not have any measurable influence on the time averaged flow field parameters [2]. the effect of this phenomenon must be investigated quantitatively to assess the induced stress in the blade. 74 vol. 20/2018 aerodynamic sources of non-stationary force loads on turbine profiles figure 7. distribution of dimensionless tangential force on the blade, points in the circle correspond to figure 5. in the figure 6, there is a scheme which describes force effects due to the unstarted supersonic flow. it is apparent, that the flow will induce torque stress in the blade. figure 7 shows distribution of the dimension-less tangential forces that affect the blade during the transition from unstarted to started supersonic flow. the torque stressing is also cyclic during the whole start up process. because the phenomenon is unstable, even small variations in the operating conditions will result in quite large changes in the stress of the blade. 4. flow separation the main cause of flow separation is adverse pressure gradient which induces back flow in the boundary layer. the boundary of the separation and position of the reattachment point are unsteady and are sources of pressure pulsation in the flow field. the boundary layer gains thickness and the streamlines detach from the surface at the point of zero velocity gradient. it is possible that further downstream the pressure gradient becomes more favourable and the flow reattaches to the surface. in this case the separation area is closed. negative incidence angle can have very similar effect, the separation forms on the opposite side of the blade. 4.1. separation at the root the root of the blade is much more resistant to the flow separation. the cascade has a very low pitch to chord ratio hence the density of the profiles in the cascade is high. relatively long interblade channel and typical profile shaping provides the root section with a wide range of operational conditions in which there is no flow separation. the experimental investigation showed, that root section cascade will maintain unseparated flow field in the range of inlet angles ι ∈ (−20◦, 30◦). the separation at the root does decrease the mass flow in the interblade channel by limiting the flow through area. the shift of effective shape of the profile can change the position and shape of the sonic line in the cascade. figure 8. flow separation on the pressure side of the blade with negative angle of incidence. (a) . ι = 0◦. (b) . ι = 20◦. (c) . ι = 30◦. figure 9. flow field in the root section cascade at different incidence angles, m2is ∼ 1.5. 4.2. separation in the tip region the experimental investigation of the tip cascades does not allow to assess effects of varying incidence angles. the models have very high aspect ratio and the blades are thin. excessive force effects induced by positive incidence angles will endanger the model and the measuring equipment. negative incidence angles will lead to very low loads and increase of flow velocity in the cascade, possibly resulting in unstarted supersonic flow. the investigation can be conducted in the region of ι ∈ (−3◦, 3◦). therefore, investigation of the tip section is limited, and the separation phenomena are not very well mapped. measurements conducted by šafařík [3] on an older last stage blade with more robust design in the tip region had the minimal inlet angle ι = −6.5◦. 75 t. radnic, m. luxa, d. šimurda acta polytechnica ctu proceedings figure 10. section near the tip, ι = −6.5◦, m2is = 1.51. figure 11. section near the tip, ι = −6.5◦, m2is = 0.6. the measurement showed that in the area behind blade’s maximal thickness there is closed area of separation, see figure 10. this separation is caused by relatively high inlet speeds and by interaction of the inlet shock wave with the boundary layer on the adjacent blade. this separation was not present during measurements in the subsonic regions, see figure 11. 5. swirl line at the trailing edge one of important aerodynamic sources of unsteady flow fields, which usually are the dominant reason of unsteady loading of blades, is a swirl line arising at the trailing edge in the near wake [4–6]. the swirl line downstream one of the blades affects the pressure distribution along a part of a suction side of the adjacent blade. this phenomenon is more noticeable in section near the tip of long blade. there, the unsteady pressure field affects (due to the geometrical arrangement of the profile cascade) a very large portion of pressure distribution on the suction side of the adjacent profile. the formation of the swirl line downstream the figure 12. tip section consisting of profiles with the tailored trailing edge, nominal condition of flow m2is = 1.68 [4]. trailing edge is characterized by the strouhal number: sh = f ·d u (1) and by reynolds number, for swirl line: re = u ·d ν . (2) the characteristic length d in case of swirl line developing at the trailing edge is the diameter of the trailing edge added to the height of the boundary layer. the criteria for evaluation of the swirl line are dependent on wide scale of flow field parameters. in a transonic flow field, the supersonic parts of the flow field in the profile cascade are subordinate to the hyperbolical description of the flow. the supersonic expansion both in the interblade channel and at the trailing edge can very favourably affect the effective dimension of the near wake, which is responsible for the swirl line origin, see figure 12. the area of confluence at the end of the near wake, has very small characteristic length. it is mainly caused by the large supersonic expansion on the sharp edges of the tailored trailing edge. the swirl line downstream the profile is insignificant and the pressure distribution along the suction side of the adjacent profile is stable. the situation in the same cascade operating at off design conditions, when the flow field is significantly more subsonic, is completely different. the supersonic expansion at the trailing edge vanishes, and the decisive size for wake and swirl line formation is the real geometric dimension of the trailing edge. the very significant swirl line takes place downstream in the wake. the individual vortices float downstream the trailing edge and periodically affect the pressure distribution along the suction side of the adjacent profile, see figure 13. 76 vol. 20/2018 aerodynamic sources of non-stationary force loads on turbine profiles figure 13. tip section consisting of profiles with the tailored trailing edge, off-design condition of flow m2is = 0.912 [4]. application of this kind of profiles in the tip region of a long rotor blade looks promising, while the aerodynamic parameters of the profile in the region of design flow conditions are good. unfortunately, it is a problem to reach the design conditions. during the start-up process the steam turbine cannot surmount the dangerous off design regimes without seriously damaging the machine. 6. conclusion the paper deals with selected undesirable phenomena present in a blade cascade. analysis of unstarted supersonic flow, flow separation and swirl line at the trailing edge of the blade had been conducted based on long term experimental research. the unstarted supersonic flow that is present in the design conditions proved to be unstable and transient. force effects, especially torque, has been identified to affect the blade when unstarted supersonic flow occurs. the flow separation does occur in non-design operating conditions, during start up, stopping or limited power output operation of the turbine. the tip region proved much more prone to detachment of the flow than the root. incidence angle deviation ∆ι = 3◦ causes separation at the tip of the blade, possibly destroying the blade cascade. the root of the blade is quite resistant to the separation which occurred at ∆ι = 20◦ and ∆ι = 30◦. the separation causes higher losses, sonic plane shift, pressure pulsations and other effects. the swirl line at the trailing edge can be caused by its round shape, which cannot be avoided for technological and structural reasons. it can induce significant pressure pulsations in the flow field. investigation of the off design transonic flow fields in blade cascades is vital for development of turbines able to standardly operate at partial power output and that can survive frequent shutdowns and start-ups in the modern intelligent power grids. list of symbols a area [m2] d characteristic length [m] f frequency of vortex shedding [hz] f force [n] m mach number [1] p pressure [pa] re reynolds number [1] st strouhal number [1] u velocity [ms−1] ι incidence angle [◦] ν kinematic viscosity [m2s−1] indexes: 1 in front of the cascade 2 behind the cascade is isentropic max maximal acknowledgements acknowledgements of the authors belong to the technology agency of the czech republic, which supported this research under grant no. th02020057. an institutional support rvo61388998 is also gratefully acknowledged. special thanks are due to the doosan škoda power co. ltd. for making this research possible. references [1] r. dvořák, p. šafařík, m. luxa, d. simurda. optimizing the tip section profiles of a steam turbine blading. in asme 2013 turbine blade tip technical symposium and course week, hamburg. 2013. doi:10.1115/tbts2013-2027. [2] m. luxa, j. příhoda, d. šimurda, et al. investigation of the compressible flow through the tip-section turbine blade cascade with supersonic inlet. journal of thermal science 25(2):138–144, 2016. doi:10.1007/s11630-016-0844-0. [3] m. luxa, et al. images of flow fields in turbine profile cascades i, ii, research report it cas no.: z-1382/06 (in czech), 2006. [4] p. šafařík. measurements on cascade tr-n-1, internal report, it csas (in czech), 1981. [5] p. šidlof, m. štěpán, v. vlček, et al. flow past a self-oscillating airfoil with two degrees of freedom: measurements and simulations. in european physical journal web of conferences, vol. 67 of european physical journal web of conferences, p. 02108. 2014. doi:10.1051/epjconf/20146702108. [6] v. vlček, j. horáček, m. luxa, j. veselý. visualization of unsteady flow around a vibrating profile. in 9th international conference on flow-induced vibration, prague. 2008. 77 http://dx.doi.org/10.1115/tbts2013-2027 http://dx.doi.org/10.1007/s11630-016-0844-0 http://dx.doi.org/10.1051/epjconf/20146702108 acta polytechnica ctu proceedings 20:73–77, 2018 1 introduction 2 experiments 3 unstarted supersonic flow 4 flow separation 4.1 separation at the root 4.2 separation in the tip region 5 swirl line at the trailing edge 6 conclusion list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0049 acta polytechnica ctu proceedings 11:49–52, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app factors influencing the fee for use of the railway infrastructure petr novák∗, václav cempírek jan perner transport faculty, studentská 95, pardubice, czech republic ∗ corresponding author: petr.novak2@student.upce.cz abstract. it is necessary to update the current system of levying charges for railway infrastructure to reflect the current market situation, to take into account the new transport strategy established by the european union, and to reflect legislative changes in relation to directive 2012/34/eu on creating a single european railway area. international trends aim to categorize tracks according to technical parameters and parameters based on shipment time, as well as many others. the existence of different opinions on the amount and way to charge fees for railway infrastructure have led the authors of this academic paper to create a plan for a system of charges to be derived from the current production factors as well as from a number of specific new factors used in countries neighboring the czech republic. keywords: prices, product factors, specific factors, transport routes, carriers. 1. introduction as defined in the white paper, "roadmap to a single european transport area – towards a competitive and resource efficient transport system," eu transport policy currently describes the point in time when the transport system’s users will cover all the costs they create. these costs include infrastructure costs (the development and maintenance of transport routes) as well as external ones – emphasized in the principle of "the user/polluter pays," which influences the environment and growing congestion. at the same time, the european commission (ec) has set requirements for lowering greenhouse gas emissions to around 20% of the 2008 level by 2030 and 60% of the 1990 level by 2050. these results should be attained by shifting 30% of road transport to railways by the year 2030 and up to 50% of such transport by 2050. the czech republic’s (cr) priority in transport policy, in accordance with directive 2012/34/eu of the european parliament and of the council of 21 november 2012 on creating a single european railway area, is implementing a competitive railway system and thereby equalizing conditions on the transport market. the extensive railway network in the cr is in direct contrast with the actuality of its fully operational lines, on which high-speed, passenger, and freight trains run and which are able to compete with road transport at roughly one fifth the rate. according to the pareto principle, 20% of the causes create 80% of the effects, which applies here in an exemplary way: the majority of railway system tracks have economic problems, because the revenues from transport do not cover the costs. 2. the current state of charging for railway infrastructure in the cr the amount to be charged is dependent on the distance traveled and the gross weight of the train, which the carrier reports to the infrastructure operator. these two parameters take into account the costs for transporting the train along the appropriate line (train-km – train kilometers) and determine the wear on the transport infrastructure caused by the train’s passage (gtk – gross ton-kilometer). charging for railway infrastructure is divided into two components. the first component used for charging is the actual use of the railway infrastructure; the second component is the fee for allocating capacity. 2.1. charging for railway infrastructure use in the cr the network statement and the ministry of finance’s current price bulletin provide annual updates on the prices for train passage on the nationwide and regional lines of the domestic railway infrastructure operated by the railway infrastructure administration (ria) in addition to the conditions for their use. these prices include not only the individual journey of a train along a track but also the organization of track-based modes of transportation, telecommunication connections between the railway operator’s employees and the train’s staff, publishing regulations, and transmitting information to carriers. the basic price is the price calculated according to a costing model using the unit prices set for passenger trains or freight trains. the basic price calculation has been set as the distance traveled by a given train along a category of track (le, lc, or lr) using data on the length and category of the sections (borders). 49 http://dx.doi.org/10.14311/app.2017.11.0049 http://ojs.cvut.cz/ojs/index.php/app petr novák, václav cempírek acta polytechnica ctu proceedings for calculating prices, the actual route covered by the train is decisive except in cases when the train is rerouted on account of a traffic disruption. the basic price is calculated according to equations 1, 2, and 3 below. cz = c1 + c2[czk] (1) where: • cz is the overall basic price for using railway infrastructure for the train’s journey [czk], • c1 is the price for using railway infrastructure for the train’s journey in the area of performance measured by the train kilometers traveled [czk], and • c2 is the price for using railway infrastructure for the train’s journey in the area of performance measured by the gross train kilometers traveled. c1 = s1e ∗ le + s1c ∗ lc + s1r ∗ lr[czk] (2) where: • s1 is the price for 1 km of a train journey (train-km) along e, c, or r category tracks [czk * km−1] and • le, lc, and lr are the distances traveled by a train along e, c, or r category tracks [km]. c2 = s2e ∗ q ∗ le + s2c ∗ q ∗ lc + s2r ∗ q ∗ lr[czk] (3) where: • s2 is the price per 1000 gross ton kilometers transported along e, c, or r category tracks [czk*gtk−1 ], • q is one thousandth of the gross weight of the train in tons. the train’s weight is the sum of the weight of all the train cars including the weight of the passengers or freight [thousandths of gross tons], and • le, lc, and lr are the distances traveled by a train along e, c, or r category tracks [km]. the price c2 is calculated separately for each section of track traveled by the given train after a change in its weight (adding/subtracting train cars, a change in the type of train from a passenger train to a train with a trainset or the reverse). 2.2. charges for capacity allocation this fee is established in conjunction with: • the length of the time interval between when the request for railway infrastructure capacity allocation was submitted and the day that has been requested for its use, • the relationship between the submitted request for railway infrastructure capacity allocation and the deadline for composing the annual railway timetable (rt) or changes planned to the timetable, and • the difficulty of processing the request. included in the charge for allocating railway infrastructure capacity is thus compensation for the process of allocating line capacity, compensation for processing the timetable for the train assigned to the given carrier request (except for printing costs and distribution aids), compensation for the train’s operative deployment, and a fee for negotiating and executing the request on short notice. the price for allocating railway infrastructure capacity is calculated according to the equation: cp k = k1 + k2 ∗ l + k3 ∗ d[czk] (4) where: • cp k is the price for allocating transport infrastructure [czk], • k1 is the rate for processing and determining the rt and allocating transport infrastructure [czk], • k2 is the rate for the construction of the train route [czk*km−1], • l is the distance of the allocated route between the starting and ending point of the track on the railway network, where the ria fills the role of the line operator or allocator of capacity [km], • k3 is the daily rate for train route allocation [czk*day−1], and • d is the number of days of the journey for which the respective route has been allocated [days]. 3. international charges for railway infrastructure in all countries, the basic system for determining railway infrastructure prices is founded on a similar basis, i.e., according to train kilometers, gross ton kilometers, or using both these transport functions on a specific track. in specific countries, other parameters that modify the final price for using railway infrastructure have also been established. these often have decisive influence on the choice of transport, and it is in the interest of each infrastructure manager to be flexible when adjusting the charges for railway infrastructure use. using an extensive analysis of the fundamentals and procedures for calculating prices for using railway infrastructure employed by infrastructure managers who proceed in accordance with directive 2012/34/eu, the authors have formed the opinion that it is necessary to supplement the current method for calculating prices with a number of specific factors. 3.1. consolidating the prices for passenger and freight transport the motivation for consolidating the basic prices for passenger and freight transport is primarily to fairly reflect the costs included in the price for using railway infrastructure. the factor of passenger/freight 50 vol. 11/2017 factors influencing the fee transport only makes it possible to reflect transport’s negative effects in a general way (e.g., the degree of wear on the infrastructure, the use of transport line capacity, and limitations on the quality of services to other carriers). 3.2. combined transport the factor of combined transport should be maintained in the new price concept for railway infrastructure use. the reason for preserving it is that it is less costly for the infrastructure manager in comparison with other freight transport and does less to limit the quality of service for railway line use by other carriers if: • trains show a high level of transport regularity and speed and • there is relatively less wear on the infrastructure with respect to a smaller volume of train wagon repair. 3.3. shipping individual wagon loads the factor of individual, fully loaded wagons should be modified to more accurately reflect the areas where this type of transport is not sufficiently competitive from the ria’s perspective. the current method of conferring timetable advantage in a blanket manner should be developed so that support for this area focuses on the sections of the first and last miles. railway competitiveness is low for precisely these sections. 3.4. track categorization in the cr, railway tracks are divided into three categories: select nationwide tracks of european importance – e, other nationwide tracks – c, and regional tracks – r. this categorization has been enacted according to ria track routing, the tracks’ importance, and the scope of transport. in the authors’ opinion, this division does not take into account the actual quality of the tracks. the benefit of dividing the tracks into more categories is to take into account their actual quality from the perspective of maximum track speed. moreover, for tracks with low track speed, this would lower the price for using the railway infrastructure, which could lead to a higher level of transport performance. 3.5. infrastructure wear this specific key factor is used to allocate costs connected to infrastructure deterioration on individual trains in conjunction with european legislation. the reason for implementing this specific factor is to take into account the different levels of wear on railway infrastructure by individual trains; so far, this has been employed using the component of price derived from gross ton kilometers. a reflection of railway infrastructure deterioration should be maintained in the price for using railway infrastructure. 3.6. etcs equipping trains with the european train control system (etcs) is key for ensuring that european railways are interoperable. regulating this specific factor reflects the gradual growth in etcs use and the overall carrier costs for supplemental train equipment. regulating this specific factor within tariff policy should derive from the condition that it have a neutral effect on infrastructure managers. thus, the advantages for trains with etcs and disadvantages for trains without etcs should be balanced. 3.7. express trains priority is given to express trains in transit to the disadvantage of other types of trains. it is appropriate to reflect the difference in the quality of services for using the railway infrastructure for these trains by an additional fee on top of the basic price – primarily for passenger trains. the reason for implementing this factor is to take into account the higher quality of services for express trains using railway infrastructure. they have priority over other trains, and their journey is therefore faster and has less interruptions – without the necessity of stopping and avoiding other trains. 3.8. the traction of traction units this specific factor takes into account the cost to the public caused by transport. it is desirable to adequately regulate this factor in comparison with other modes of transportation and for it to simultaneously motivate carriers to lower the cost to the public. protecting the environment against the negative effects of transport is one of the ec’s key strategic goals. however, the authors believe that measures taken in this area for railway transport are not distinctly restrictive in nature – conversely, they tend to support their use. 3.9. reducing noise levels taking into account the noise levels of trains and other means of transport requires composing comprehensive methods to make it possible to quantify the public costs caused by individual modes of transport. for this reason, it is appropriate to implement the factor of "reducing noise levels" once there is a unified and time-tested european method that is able to take into account roadway noise levels. 3.10. the degree of track utilization the general differentiation of peak and off-peak periods makes it possible to optimize the use of railway infrastructure capacity. the factor of utilization is appropriate only for certain tracks, and it must always be established on an individual basis. support for optimizing railway infrastructure utilization aims to decrease the overloading of individual 51 petr novák, václav cempírek acta polytechnica ctu proceedings routes primarily during peak hours as well as to increase the use of tracks that currently have a distinct reserve capacity. international experience shows that it is not possible to influence certain effects of transport using universal tools; rather, it is necessary to react to them individually, e.g., by taking into account the operational peak and off-peak times for individual tracks. 3.11. a train’s minimum power density according to the authors, one specific factor that limits excessive consumption of railway infrastructure capacity is caused by insufficient traction unit performance and the consequently low speeds of certain trains. in turn, this lowers the risk of events occurring that block infrastructure capacity – caused by the gridlock of trains with insufficient traction units. 4. conclusion charging for roadway and railway infrastructure is a sensitive topic, because any type of change has effects on financing infrastructure. when looking for an optimal solution for charging fees, it should ensure the equalization of transport between branches and a competitive transport system; it is also necessary to take into account the policy of both the cr and the eu. analysis of the factors listed above will be the basis for establishing a new method for calculating railway infrastructure prices. for a comprehensive evaluation, it is necessary to ascertain the relevant information and the czech ministry of transport’s projected approach, which is tied to fulfilling the goals of the czech republic’s transportation policy and incorporating eu directives and regulations. in comparing the cr’s railway and roadway transport, where charges are levied for only 2.6% of the length of the roadway network and 98% of the railway network, it can be seen that there is obvious discrimination against railway transport in inter-branch competition. 52 acta polytechnica ctu proceedings 11:49–52, 2017 1 introduction 2 the current state of charging for railway infrastructure in the cr 2.1 charging for railway infrastructure use in the cr 2.2 charges for capacity allocation 3 international charges for railway infrastructure 3.1 consolidating the prices for passenger and freight transport 3.2 combined transport 3.3 shipping individual wagon loads 3.4 track categorization 3.5 infrastructure wear 3.6 etcs 3.7 express trains 3.8 the traction of traction units 3.9 reducing noise levels 3.10 the degree of track utilization 3.11 a train's minimum power density 4 conclusion acta polytechnica ctu proceedings doi:10.14311/app.2017.8.0027 acta polytechnica ctu proceedings 8:27–29, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app the current view on the use of reconstruction materials in dentistry radka vrbovaa, ∗, pavel bradnaa, martin bartosa, lucie himmlovaa, tomas horazdovskyb a institute of dental medicine, first faculty of medicine, charles university and general university hospital in prague, prague, czech republic b department of physics, faculty of mechanical engineering, czech technical university in prague, prague, czech republic ∗ corresponding author: vrbova@vus.cz abstract. the hardest tissue in the human body is the enamel which covers the anatomical crowns of teeth. it must be resistant to mechanical stress and the chemical attack of many substances from food, drinks and products of the metabolism of bacteria present in the oral cavity. these low ph substances dissolve the mineral components of enamel, cause tooth demineralization, and lead to decay or erosion damage with the irreversible loss of dental hard tissues and the necessity of their reconstruction. the range of dental materials intended for dental tissue reconstruction is extensive. dental amalgam can be mechanically applied into the strongly stressed lateral segments of teeth. the use of amalgam is, however, in decline, with the possible health risks attributed to it, coupled with the need to extensively prepare tooth tissue promoting a shift towards using aesthetically and biologically favourable dental ceramic and polymeric materials instead. current developments also concentrate on these materials to reinforce this, with polymeric composite materials based on methacrylates with varying amounts of inorganic fillers at the forefront. these materials are distinguished by their good mechanical and aesthetic properties and wear resistance. however, polymerization shrinkage and a strong hydrophobic nature does not allow for their direct bonding to hard dental tissues. risks associated with the release of residual free monomers from the structure to the environment, which may cause health complications, mainly allergic reactions in sensitive individuals, have been monitored recently. further development in the field of composite materials aims to reduce or completely eliminate these negatives. keywords: hard dental tissues, tooth decay, erosion, dental materials. 1. introduction teeth are specific formations in the oral cavity which are mainly composed from hard tissues such as enamel, dentine and cementum. enamel covers the surface of anatomical crowns and is the hardest tissue in the human body due to the high content of mineral substances. the so-called prismatic structure is typical for enamel, with closely spaced prisms bound with interprismatic substance (fig. 1a). enamel prisms with a diameter of 4-8 µms are formed predominantly from hydroxyapatite nanocrystals. the interprismatic mass also contains proteins, except hydroxyapatite, which serve as a binder for the individual prisms. dentine is mineralized less than enamel, is softer, and contains about 20-30 vol. % of proteins and 20-30 vol. % of water, and is biologically active compared to enamel. typical structural elements are known as dentinal tubules, canals in the dentine where extensions of the odontoblasts exist alongside dentinal fluid (fig. 1b). figure 1. a) prismatic structure of enamel after phosphoric acid etching (cross-section); b) the surface of dentin with dentinal tubules after phosphoric acid etching (cross-section). (sem, jsm5500-lv, jeol, tokyo, japan). figure 2. a) 3d model of the damaged tooth by caries from micro-ct (skyscan 1272, bruker, belgium); b) the longitudinal cross section of a tooth with tooth decay in the enamel and dentin (optical microscope olympus, smz 2mt, japan). 27 http://dx.doi.org/10.14311/app.2017.8.0027 http://ojs.cvut.cz/ojs/index.php/app r. vrbova, p. bradna, m. bartos et al. acta polytechnica ctu proceedings figure 3. the surface of enamel exposed to drinks for 12 hours: a) intact enamel; b) coca-cola, ph 2.5; c) fanta, ph 2.8; d) sprite ph 2.7. (sem, jsm5500-lv, jeol, tokyo, japan). 2. the typical damages of hard dental tissues 2.1. tooth decay due to the effect of products of the metabolic conversions of cariogenic bacteria present in dental plaque, there is a local demineralization in the subsurface layers of enamel. if the process reaches the dentineenamel border, it will break the surface layers of the enamel and will lead to the formation of a cavity (fig. 2a, b). 2.2. erosion in the case of dental tissue erosion, the demineralization is of non-bacterial origin. it results from the attack of substances with high acidity and simple sugars content. from this point of view, so-called "soft drinks" are very aggressive with a low ph, containing e.g. phosphoric acid or citric acid (fig. 3 b, c, d). they cause the dissolution of enamel mineral components; reduce its surface hardness and resistance to abrasion or mechanical stress. the result of the above-mentioned conditions may be the loss of dental hard tissues with the necessity of their reconstruction, which is often complicated and costly. additionally, the lifetimes of reconstructions are always limited. 3. polymeric materials for the reconstruction of hard dental tissues the range of dental materials intended for the reconstruction of lost or damaged dental tissues, whether as a result of carious processes, extensive erosions or traumatic damages is very broad. the most frequently used are composites, whose matrix consists mainly of methacrylate polymers (e.g. bis-gma, tegdma, udma, bis-ema) and the fillers are ba, sr, zr ground glass with surfaces treated by silane coupling agents and amorphous sio2 added primarily for adjusting the thixotropic properties of materials and pigments. these are mainly light-cured materials containing photoinitiators of polymerization reaction, usually camphorquinone (2,3-bornandion), propanedione (1phenyl-1,2-propanedione) or lucirin tpo (biphenyl (2,4,6-trimetylbenzoyl-diphenylphosphine oxide), often in conjunction with tertiary amine co-initiators. figure 4. a) the adhesive material leaking into the dentinal tubules; b) damaged marginal seal between the filling material and enamel. (sem, jsm5500-lv, jeol, tokyo, japan). they must also contain inhibitors of the polymerization reaction. the hardening of materials occurs by radical polymerisation, when molecules of photoinitiators absorb photons of light from curing lamps forming free radicals, which initiate very fast polymerization reactions. mainly halogen or led curing lamps emitting light in the wavelength range 400 to 515 nms are currently used for curing of materials. it is very important that the spectrum of emitted light from the curing light must overlap the absorption spectrum of the photoinitiator present in material to ensure the best possible polymerization [1, 2]. composite materials are characterized by very good mechanical and aesthetic properties, and are indicated for lateral and front sections of the dental arches. however, these materials are hydrophobic and do not adhere to tooth structure comprising water. therefore, quality micromechanical binding provides adhesive systems that flow into the previously etched microrelief of dental tissue (fig. 4a). adhesive systems include hydrophilic monomers with the ability to wet moist surfaces and bond them with hydrophobic monomers which are chemically combined with a composite material after polymerization. the main disadvantage of these materials is their shrinkage due to polymerization. this may cause damage to the marginal seal, i.e. loss of contact between the dental tissue and the filling material (fig. 4b). the polymer materials also show a possible risk of residual free monomers releasing into the surrounding environment, which may give rise to health problems in susceptible individuals, especially to allergic reactions [3–6]. 28 vol. 8/2017 the current view on the use of reconstruction materials in dentistry 4. conclusion further development of polymer composites is driven by efforts to improve their mechanical and aesthetic properties, increase wear resistance, reduce polymerization shrinkage and achieve significant improvements in biocompatibility. acknowledgements this work was supported by grant prvouk p28/lf1/6 mšmt. references [1] p. bradna, l. comba, j. froněk, d. houšová. kam směřuje vývoj polymeračních lamp? česká stomatologie 110(3):61–68, 2010. [2] j. froněk. studium vlivu vlastností polymeračních lamp na vytvrzení kompozitních materiálů s různými fotoiniciačními systémy. master’s thesis. [3] m. syed, r. chopra, v. sachdev. allergic reactions to dental materials-a systematic review. journal of clinical and diagnostic research 9(10):ze04 – ze09, 2015. doi:10.7860/jcdr/2015/15640.6589. [4] i. stoeva, a. kisselova, m. zekova. allergic contact stomatitid from bisphenol-a-glycidyldimethacrylate during application of composite restorations. a case report. journal of imab annual proceeding (scientific papers) pp. 45 – 46, 2008. book 2. [5] s. gupta, p. saxena, v. pant, a. pant. release and toxicity of dental resin composite. toxicology international 19(3):225–234, 2012. doi:10.4103/0971-6580.103652. [6] m. goldberg. in vitro and in vivo studies on the toxicity of dental resin components: a review. clinical oral investigations 12(1):1–8, 2008. doi:10.1007/s00784-007-0162-8. 29 http://dx.doi.org/10.7860/jcdr/2015/15640.6589 http://dx.doi.org/10.4103/0971-6580.103652 http://dx.doi.org/10.1007/s00784-007-0162-8 acta polytechnica ctu proceedings 8:27–29, 2017 1 introduction 2 the typical damages of hard dental tissues 2.1 tooth decay 2.2 erosion 3 polymeric materials for the reconstruction of hard dental tissues 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0049 acta polytechnica ctu proceedings 13:49–54, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app plasma treatment impact on physical and chemical properties of polymeric fibers radim hlůžeka, ∗, zdeněk prošeka, b, jan trejbala, josef fládra, štěpán potockýc a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b university centre for energy efficient buildings of technical university in prague, třinecká 1024, 273 43 buštěhrad, czech republic c czech academy of science, institute of physics, na slovance 1999/2, 182 21 prague 8, czech republic ∗ corresponding author: radim.hluzek@fsv.cvut.cz abstract. presented work focuses on chemical and physical properties of plasma modified polymeric macro-fibers. polyethylene terephthalate (pet) and polypropylene (pp) fibers having approx. 300 µm in diameter were modified using cold oxygen plasma in order to achieve their surface changes needed for durable bond and adhesion with cement matrixes. a duration of plasma modification differed between 5 to 480 seconds, where an effect of the treatment was examined. fiber surfaces chemical changes were researched via wettability measurement with demineralized water (the measurement was repeated immediately and after 1, 7 and 30 days to find out the changes stability). physical changes were studied by means of weight balance (determination of weight loss) and tensile strength tests. it was found that wettability was enhanced significantly – up to two times, while mechanical properties of treated fibers decreased only slightly. keywords: polymer macro-fibers, plasma treatment, wettability, tensile strength. 1. introduction usage of man-made macro fibers as reinforcement of any composites become a commonality in majority of industry fields, civil engineering including. in the final consequence, a small bulk property improvement has strong impact in civil engineering due to mass production [1]. fiber-reinforced materials exhibit good mechanical properties, e.g. high mechanical resistance (abrasion, impact resistance), ductility, water resistance etc. [2, 3]. the main task of the macro fiber addition is the distribution of shrinkage into several small cracks in the case of materials based on shrinking binders (mainly lime and cement), and to prevent the single crack openings after linear elastic response in the case of loaded samples [1]. in the civil engineering, the macro fiber reinforcement is the most often used for a) production of watertight concretes and concretes exposed to risk of steel reinforcement corrosion (absence of cracks disallows water penetration), and b) for production of large-scale construction exposed to temperature or moisture changes, dynamic loads, point loads (the fiber reinforcement provides a compactness – hence usability – of materials after the crossing the material loading capacity) [4, 5]. the fiber reinforcement can be classified by fibers material, diameter, tensile strength and their modulus of elasticity. as material, polymeric fibers can be used. both polymeric (in particular pet and pp) fibers have high tensile strength equal to about hundred or even thousands mpa. their diameter is equal to tens or hundreds micrometers. other characteristic property is low ratio of diameter to length (and related high specific surface enabling better stress transferring from matrix to fibers) and favorable cost [6, 7]. polymeric fibers reveal low surface wettability (hydrophobicity). on the other hand, fiber reinforcement requires good adhesion between the fiber surface and matrix. to improve the mechanical strength of reinforced materials, adhesion between the fiber surfaces and matrix must be unambiguously ensured [5, 8]. the required adhesion can be modified by fiber surface treatment by chemical and physical methods [7, 9– 13]. currently, the newly introduced plasma treatment becomes popular as a progressive physico-chemical method. the low-pressure plasma treatment represents a universal, efficient and eco-friendly alternative for surface modifications. plasma can be defined as ionized gas (composed of electrons, ions, and neutral species). the mechanism for plasma surface modification relays on surface atoms replacement by oxygen atoms and formation of polar groups. the presence of polar or functional chemical groups enhances the reactivity with the matrix based on cement or lime binder (both contain water) [5]. in the present work, we report on the modification of polyethylene terephthalate (pet) and polypropylene (pp) fibers by oxygen plasma treatment. the influence of the plasma treatment on the contact angle, the weight and the tensile strength of the fibers is studied. 49 http://dx.doi.org/10.14311/app.2017.13.0049 http://ojs.cvut.cz/ojs/index.php/app r. hlůžek, z. prošek, j. trejbal et al. acta polytechnica ctu proceedings figure 1. digital camera photo of pet (left), pp fibers (right). 2. materials 2.1. polymer macro-fibers two different types of fibers were used: pet and pp. both of them were made by spokar manufacturer (czech republic). they were primary made for the production of most types of brushes (tooth, paint etc.) and brooms [14]. the fibers had diameter equal to approx. 300 µm (measured using mechanical micrometer), original length was 1200 mm (subsequently chopped according to need). an industrial water flushable sizing is standardly applied onto fiber surfaces as an integral part of their production. fibers are shown in figure 1. 3. experimental methods 3.1. plasma treatment to improve the wettability and related adhesion between fiber surfaces and surrounded matrix, oxygen treatment in inductively coupled plasma was realized using device tesla vt 214 (13.56 mhz). parameters of the whole process differing only in exposition time from 5 to 480 seconds, the others were constant; total power of rf source 100 w, total gas pressure 20 pa, oxygen flow 50 sccm. 3.2. contact angle measurement according to reference [15], using horizontal direct optical method, contact angles were measured between meniscus of distilled water (adhering on fiber surfaces) and fiber perpendicularly submerged into the water. thus obtained results were averaged from 6 independent measurements. 3.3. weight ratio analysis due to ion bombardments and polymer oxidation and surface structuring (etching) during the plasma treatment, the fibers can be damaged. weight ratio observation carried out before and after treatment enables to assess the fiber weight loss, which is connected to tensile strength capacity loss. in order to quantify the fiber weight loss, fibers were weighed before and immediately after plasma treatment on weighing-machine kern alj 120-4 (d = 0.1 mg). 3.4. tensile strength tests physical and chemical processes mentioned earlier can influence the fibers tensile strength. to determine the extent of fibers mechanical properties damage, tensile strength test was carried out using loading frame web tiv ravestein fp100. the test was displacement controlled at a constant rate equal to 6 mm/min in the case of pet and 3 mm/min in the case of pp fibers, respectively. the results were averaged from 6 independents experiments for each plasma treatment time of both fiber types. 4. results and discussions 4.1. contact angle measurement contact angle measurement revealed 84.7±2.2◦ on untreated pet fibers and 88.3±0.6◦ on untreated pp fibers. the results provided on the treated fibers showed the unexceptionable positive plasma impact of the fibers wettability. all treated fibers independently on plasma exposition time showed significant decrease of contact angle. the angle measured immediately after plasma treatment on both types of fibers was moving between approx. 25–30◦. after 1, 7 and 30 days from the treatment, contact angle increased up to 65–70◦ for pet fibers and for the pp fibers up to 55–70◦. besides, the results showed that it is not effective to do any treatment longer duration than several seconds. another finding is the fact that storage of fibers, even for one day, is very inefficient, and the contact angle increase rapidly towards the reference values. dependence between contact angle, duration of plasma treatment and time of storage in atmospheric conditions is shown in figure 2 and figure 3. based on these finding, it was decided that plasma treatment duration realized within other experiments throughout the study were shortened to 120 seconds only. demonstration of the fibers wettability captured during contact angle measurement is shown in figure 4. 4.2. weight ratio analysis weight ratio observation performed on fibers before and after plasma treatment showed that the weight 50 vol. 13/2017 physical and chemical properties of polymeric macro-fibers figure 2. pet fibers – dependence between contact angle, duration of plasma treatment and time of exposure to atmospheric conditions. figure 3. pp fibers – dependence between contact angle, duration of plasma treatment and time of exposure to atmospheric conditions loss is negligible. fibers treated by 120 seconds exhibited practically no weight loss, just less than 0.25 %. comparing both types of fibers, it can be said that the weight loss is the same until 30 seconds of plasma treatment. after 30 seconds of treatment, there are some small difference between pet and pp fibers. after the longest time of treatment (120 seconds), the maximum averaged weight loss was 0.06 % and 0.21 % for pet and pp fibers, respectively. based on these findings, it can be assumed that mechanical properties of treated fiber were not affected by plasma treatment, which is a positive result. graphic relation between weight loss and duration of plasma treatment is shown in figure 5. 51 r. hlůžek, z. prošek, j. trejbal et al. acta polytechnica ctu proceedings figure 4. contact angles captured during the wettability measurement figure 5. relation between fibers weight loss and duration of plasma treatment obtained before and after plasma treatment. figure 6. dependence between fibers tensile strength and duration of plasma treatment. 52 vol. 13/2017 physical and chemical properties of polymeric macro-fibers 4.3. tensile strength tests tensile strength (more precisely load carrying capacity) of reference and plasma treated fibers was approximately constant for all plasma treatment times (measured on reference, 10, 60 and 120 seconds treated fibers). pet fibers reached on approx. 20 n, while pp fibers on approx. 32 n. as it was already shown by weight loss ratio mentioned earlier, fiber tensile strength test confirmed that the plasma treatment had no significant effect on fiber mechanical performance. the relation between fibers tensile strength and duration of plasma treatment is shown in figure 6. 5. conclusions polypropylene (pp) and polyethylene terephthalate (pet) commercial macro fibers spokar having approx. 300 µm in diameter were surface treated by mean of low-pressure cool oxygen plasma treatment in order to achieve their chemical and physical surface changes. the aim of the executed treatment was to improve fiber surface wettability (from hydrophobic to hydrophilic) and morphology (from smooth to roughened), both standardly required for strong interphase interaction with cement matrix, when fibers used as reinforcement. the plasma modification differed in duration (from 5 to 480 seconds). an effect on physico-chemical changes on fiber surfaces was examined by wettability measurement, wight loss ratio (fibers were weighed before and after plasma treatment) and finally by fiber tensile strength tests, when the main focus of the research was targeted to impact of plasma treatment on fiber mechanical performance. besides, time-stability of chemical changes on modified fibers was researched – wettability between fiber surfaces and distilled water was examined using direct optical set for contact angle measurement immediately, 1, 7 and 30 days after plasma modifications. it was found out that: • the wettability of plasma treated (regardless to duration of treatment time) both pp and pet fibers increased almost three times, if compared to reference samples. • the modified wettaility was very short-term. it was degraded almost to reference values already after one day, when fibers exposed to standard atmospheric conditions. • impact of realized treatment did not have no obvious effect on fiber mechanical performance, as demonstrated by no weight loss and no tensile strength changes for much as treated fibers. acknowledgements this work was financially supported by czech technical university in prague – sgs project sgs16/201/ohk1/3t/11, by the czech science foundation research project 15-12420s and by the ministry of education, youth and sports within national sustainability programme i, project no. lo1605. the work occurred in frame of the lnsm infrastructure (no. lm2015087). references [1] p. k. mallick. materials. in fiber-reinforced composites, materials, manufacturing and design, chap. 2. crc press, 2007. [2] j. trejbal, l. kopecký, s. potocký, j. fládr. the effect of glass fiber reinforcement on flexural strength of lime-based mortars. in 53rd conference on experimental stress analysis, pp. 450–453. czech technical university in prague, 2015. [3] k. šeps, j. vodička. fibre reinforced concrete with recycled concrete and insulation material stered. advanced materials research 1054:262–266, 2014. doi:10.4028/www.scientific.net/amr.1054.262. [4] z. lin, v. li. crack bridging in fiber reinforced cementitious composites with slip-hardening interfaces. journal of the mechanics and physics of solids 45(5):763–787, 1991. [5] j. trejbal, v. šmilauer, a. kromka, et al. wettability enhancement of polymeric and glass micro fiber reinforcement by plasma treatment. in nanocon 2015 7th international conference on nanomaterials – research and aplication – conference proceedings, pp. 315–320. tanger, 2015. [6] j. bartoš. modification of pet fiber surfaces to increase their cohesion with cement matrix. diploma thesis. czech technical university in prague, 2014. [7] j. trejbal. modification of glass fibers surface and their application in mortars used for reconstructions of historical buildings. diploma thesis. czech technical university in prague, 2014. [8] p. chen. influence of fiber wettability on the interfacial adhesion of continuous fiber-reinforced ppesk composite. journal of applied polymer science 102:2544–2551, 2006. doi:10.1002/app.24681. [9] s. elsaka. influence of chemical surface treatments on adhesion of fiber posts to composite resin core materials. dental materials 29:550–558, 2013. doi:10.1016/j.dental.2013.03.004. [10] d. cökeliler, s. erkut, j. zemek, et al. modification of glass fibers to improve reinforcement: a plasma polymerization technique. dental materials 23(3):335–342, 2007. doi:10.1016/j.dental.2006.01.023. [11] m. kim, m. kim, k. rhee, s. park. study on an oxygen plasma treatment of a basalt fiber and its effect on the interlaminar fracture property of basalt/epoxy woven composites. composites part b: engineering 42(3):499– 504, 2011. doi:10.1016/j.compositesb.2010.12.001. [12] h. wu, v. li. fiber/cement interface tailoring with plasma treatment. cement and concrete composites 21(3):205–212, 1999. doi:10.1016/s0958-9465(98)00053-5. [13] j. trejbal, j. fládr, l. kopecký. utilization of plasma treated polymeric macro-fibers as reinforcement in concrete constructions. advanced materials research 1144:70–75, 2017. doi:10.4028/www.scientific.net/amr.1144.70. 53 http://dx.doi.org/10.4028/www.scientific.net/amr.1054.262 http://dx.doi.org/10.1002/app.24681 http://dx.doi.org/10.1016/j.dental.2013.03.004 http://dx.doi.org/10.1016/j.dental.2006.01.023 http://dx.doi.org/10.1016/j.compositesb.2010.12.001 http://dx.doi.org/10.1016/s0958-9465(98)00053-5 http://dx.doi.org/10.4028/www.scientific.net/amr.1144.70 r. hlůžek, z. prošek, j. trejbal et al. acta polytechnica ctu proceedings [14] spokar a.s. fiber production. http: //www.spokar.com/cz/technologie/vyroba-vlaken. [15] j. trejbal, l. kopecký, s. potocký, z. remeš. direct optical method for contact angle measurement on micro fibers. in nanomaterials and nanotechnology in civil engineering 2015 – conference proceedings, pp. 89–90. czech technical university in prague, 2015. 54 http://www.spokar.com/cz/technologie/vyroba-vlaken http://www.spokar.com/cz/technologie/vyroba-vlaken acta polytechnica ctu proceedings 13:50–55, 2017 1 introduction 2 materials 2.1 polymer macro-fibers 3 experimental methods 3.1 plasma treatment 3.2 contact angle measurement 3.3 weight ratio analysis 3.4 tensile strength tests 4 results and discussions 4.1 contact angle measurement 4.2 weight ratio analysis 4.3 tensile strength tests 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0059 acta polytechnica ctu proceedings 5:59–62, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app the priority of international freight expresses in the overlapping section of rfc 7 and rfc 9 kolín–česká třebová přemysl šrámek∗, tatiana molková university of pardubice, jan perner transport faculty, department of transport technology and control, studentská 95, 532 10 pardubice, czech republic ∗ corresponding author: st24539@student.upce.cz abstract. this article deals with the railway traffic control and railway traffic equipment in terms of implementation of tsi. the main goal is to research the optimal priority of international freight expresses, especially in the overlapping sections, which represent critical line sections of european freight corridors. this article shows this problem in the context in the overlapping section of rfc 7 and rfc 9 kolín–česká třebová. there are mentioned possibilities of using of etcs and implementation of automatic train paths. keywords: capacity, freight expresses, freight corridors, overlapping section, periodic timetable, traffic control . 1. introduction one of the aims of european transport policy is to redirect capacity of road freight traffic to other modes of transport, where rail transport is an interesting and environmentally friendly alternative, especially in terms of speed, availability and amount of transported cargo. on the other hand, by customers is required delivery time is guaranteed by carriers, what is unfortunately not always possible, especially due to high utilization of european rail infrastructure. the solution is the construction of new lines, increasing the capacity of existing lines through construction and reconstruction measures or through operational measures. among the operational measures may be included alternative routing of trains as well as changing priorities of a particular type of train, e.g. train category fex (freight express). 2. aims the aim of this article is to find the optimal priority of international freight expresses to satisfy the customers required delivery times due to optimization of railway traffic control. unfortunately, in the czech republic is very problematic to buy up structural land for extension of railway infrastructure, e. g. for construction of another track line. the railway infrastructure of the overlapping section of rfc 7 and rfc 9 kolín – česká třebová is equipped with the new modern signalling system with etcs level 2. the possibility to improve the signalling system is only implementation etcs level 3 or implementation of automatic train paths. however, the easiest variant is to change the priority of freight expresses and observe the passenger transport impact. 3. materials and methods there are mentioned some information about european freight corridors rfc, leaflet uic 406, deadline cargo mode and the simulation method with using of simulation tool simut (average delay increment (adi) calculation). 3.1. rail freight corridors (rfc) on the basis of the european parliament and the eu council no. 913/2010 for competitive freight and no. 1316/2013, which is created the connecting europe facility (cef), is the gradual establishment of rail freight corridors rfc (rail freight corridors). the main aim of the operation of these corridors is to strengthen the competitiveness of railways synergies between rail systems and harmonization of allocation interstate freight routes by national infrastructure managers [1]. czech republic, respectively czech railway infrastructure administration (cria), as the infrastructure manager and allocator of capacity, is a member of rfc 5 (baltic – adriatic), rfc 7 (orient / east-mediterranean), rfc 8 (north sea – baltic) and rfc 9 (czechoslovak corridor). each corridor is operated on the basis of the corridor information document (cid); each corridor has a single point of contact (c oss) and allows carriers to apply international prearranged train paths facilitated freight via c-oss. the request must be filed by information system rne pcs and must include the cross-border section. this paper deals with the model study of common section of rfc 7 and rfc 9 kolín – česká třebová. 3.1.1. rfc 7 rfc 7 – the orient corridor runs from central europe to easternand southern-europe connecting 7 member states – czech republic, austria, slovak republic, 59 http://dx.doi.org/10.14311/app.2016.5.0059 http://ojs.cvut.cz/ojs/index.php/app přemysl šrámek, tatiana molková acta polytechnica ctu proceedings hungary, romania, bulgaria and greece. the total length of main lines is approx. 3 900 km and the length of alternative and connecting lines is almost 2 500 km altogether. most limiting factors of rfc 7 are low capacity, speed limit, limited length of trains, limited axle load, not electrified sections and lack of adequate safety equipment (signalling track circuits with 25 hz frequency, etcs, gsm-r, etc.). the railway infrastructure managers and capacity allocation companies responsible for establishing and running rfc 7 are committed to offer reliable, high-quality, competitive transport services in order to increase the market demand, to operate the infrastructure costeffectively on the long run through harmonization of technical and procedural conditions and to facilitate the environmentally sustainable development of the european economy and the achievement of a better quality of life for its people [2]. in capacity analysis, published in implementation plan of rfc 7 [2], there are found lines with capacity utilization higher than 90 %. the longest sections with this high capacity utilization are situated in the czech republic, concretely sections poříčany – pardubice (65 km) and choceň – česká třebová (25 km). the whole second section is in the model study, which results are published in this paper, included. the first section is included, too, without the section kolín – poříčany (20 km). 3.1.2. rfc 9 rfc 9 the czech-slovak rail freight corridor (cs corridor) runs from prague to čierna nad tisou (slovak-ukrainian border) connecting 2 member states czech republic and slovak republic. the total length of main lines is 972 km and the length of alternative and connecting lines is 276 km altogether. most limiting factors of rfc 9 are the same as in rfc 7. in capacity analysis, published in implementation plan of rfc 9 [3], there are found lines with capacity utilization higher than 90 %. 3.2. the uic 406 leaflet the uic 406 leaflet, completely overhauled in june 2013, is one of the most important documents about the capacity calculation. there is published the approach to calculate capacity consumption by compressing a timetable and to evaluate the number of possible train path to a node, line or corridor. the compression method consists of five steps – defining infrastructure and timetable boundaries, defining section for evaluation, calculation capacity consumption, evaluating capacity consumption and evaluation available capacity. corridor is in the uic 406 defined as the main international and national connection and thus usually stretches over several hundred kilometres. corridors may overlap with one another overlapping section is illustrated on the figure 1 [4]. 3.3. the deadline cargo mode periodic schedules are in rail passenger transport commonplace. in order to satisfy also the end users of figure 1. overlapping train path line sections [4] rail freight and ensure its full competitiveness, it is necessary to trace freight trains in the fixed periodic routes, which would have removed one of the main rail transport disadvantages against road transport. to increase the transport speed and reduce the overall delivery time is also desirable international freight train paths were carried out accurately under deadline cargo for this purpose it may be necessary to change the current train priority so that international freight expresses (fex) weren’t in their paths delayed by stop passenger trains. this measure in addition to increasing the speed of transport and reducing the overall delivery time also contributes to reducing the overall energy efficiency of rail freight, which will no longer need to waste power electricity due to start heavy freight trains. 3.4. the simulation method in this part is said something about the simulation method in context of the transit of international freight express trains (fex) in the deadline cargo mode in the overlapping section of rfc 7 and rfc 9 kolín česká třebová. in this section there is a total amount of 14 railway stations, in which it is possible overtaking trains (double track line). interstation sections are divided by an automatic block into track sections; all crossing safety devices are equipped with gates. for simulation was used simulation tool simut, developed by pavel krýže, phd. from cria. the simulation tool simut is developed in visual basic, therefore it can run on every pc with microsoft office. it has to be input the option of a simulation, then railway stations and their shortcuts, station tracks for each railway station, line tracks, connection of station and line tracks, length of interstation departments and amount of line departments, type of train for priority, type of train for each number of a train and the path of each train. the simulation program simut can put new paths in a daily timetable with the solution for arisen path conflicts. it was created the daily timetable, which 60 vol. 5/2016 the priority of international freight expresses variant adilong distance passenger transport regional passenger transport fex total low priority 5,40 10,17 5,12 5,87 d1 priority 7,22 19,35 -5,63 4,03 high priority 7,67 18,57 -8,49 3,14 table 1. adi for different variants of priority of fex trains. included the amount of 384 trains, concretely 160 express trains, 34 speed-up passenger trains, 46 stop passenger trains and 144 fex trains. the passenger transport trains were concentrated especially from 6 am to 23 pm. all trains were conducted in periodic timetable. within the simulation program was established the average delay increment (adi). the average delay increment was calculated by dividing the difference between total output and total input delay and the total number of trains. this indicator was calculated as an ongoing basis for each simulation run, so the total for the entire graph (all simulation runs). the indicator was also calculated for different types of transport, i.e. for long-distance passenger transport, regional passenger transport and freight transport. in order to achieve measurable results of transiting fex within rfc in the deadline cargo mode, the freight transport segment was narrowed to only those trains, for which it was created recurring schedule, counting 3 pairs of trains every hour (the period 20 min) on the prescribed speed of 100 km/h (total amount of 144 trains). in the real operation the other freight trains shouldn’t delay the trains of fex category. three variants were created, reflecting the priority of fex. the first variant corresponded more or less simultaneous the operation in the czech republic, when international freight trains of the fex category were delayed by all passenger trains including stop passenger trains. in the second variant there was set priority of trains, corresponding to the transport prescription d1 of cria [5], when the fex trains overtook stop passenger trains. and in last, the third variant, received international freight trains of the fex category the second highest priority, not just for passenger trains category ex (expresses). as part of the simulation was set for all simulation runs random entry delay based on the exponential probability distribution. there were solved conflicts of station tracks, freight trains were allowed to ride before their schedule time (in the case of free capacity). for each variant was made a total of 365 runs of simulation (for a daily timetable) [6]. 4. results and discussion depending on the other train categories (multicriteria decision model – stop passenger train has got coefficient 1) it was set low priority of fex trains (0,5), priority of trains by transport prescription d1 made figure 2. average delay increment for different variants of fex train priority. by cria (1,7) and high priority of fex trains (1,9) – results are displayed in table 1. for each run of simulation had to be done 34 500 solutions by simulation program simut, it meant in total to count for 365 runs of simulation 12 592 500 solutions for each variant for three different variants had to be counted 37 777 500 solutions of daily timetable with random entry delay based on the exponential probability distribution. the most important is the total average delay increment this indicator is getting lower with the increasing priority of fex trains. in this research we are getting back to the beginning of railways in the territory of czech republic to years, when cargo transport was preferred. in the growing amount of fex trains, which are operated on rfc, it seems quite logical to overtake some passenger transport trains by fex trains to balance the railway transport. for cria is this model convenient because of higher fee payed for using of the railway infrastructure (which is generated by fex trains). and especially the high priority of fex trains fulfil at most the terms of deadline cargo mode, which is preferred by customers. the results from table are illustrated on figure 2. to explain more the results presented on the figure 2, it must be said the adi is for fex trains negative (d1 and high priority) due to allowing of riding before their schedule time. the schedule was created and new paths were included on the basis of simultaneous operation (low priority). therefore, if the fex train could overtake in other variants the train of another category, it did it and then the average delay increment of fex was negative, on the other hand the adi 61 přemysl šrámek, tatiana molková acta polytechnica ctu proceedings of the other train category increased. from the table 1 is apparent, too, all train categories couldn’t be conducted in the overlapping section of rfc 7 and rfc 9 kolín česká třebová in periodic timetable. if will be possible to conduct in periodic timetable the fex trains, it is inappropriate to conduct the same way the other train categories due to increasing adi. the solution could be to adjust paths of stop passenger trains – e.g. incorporate some deviation of periodic timetable. in long distance passenger transport the operation could be solved by consolidation of expresses, because nowadays short passenger expresses (locomotive + 3 wagons) often cause the stopping of heavy fex trains due to overtaking. if the passenger express will have more wagons and to final directions will be divided in intermediate station, it resists more capacity for fex and stop passenger trains. in the section kolín česká třebová is this rule very good applicable – present expresses could go together in one longer express train from prague to česká třebová, where will be divided to directions brno and ostrava. in the context of implementation of tsi-taf in the overlapping section, it means operation of etcs level 2 with full-functional gsm-r, it is offered the extension to etcs level 3 or the implementation of automatic train paths on etcs level 2. both variants enable to improve the railway traffic control technology and to reduce the amount of human interventions to do the railway traffic safer and well-timed. 5. conclusions to boost the attractiveness of railway freight transport it has to be kept customers required delivery times. to keep the customers required delivery times railway freight transport has to be operated in deadline cargo mode. to operate the railway freight transport in deadline cargo mode it has to be guaranteed train path stability. to guarantee the train path stability for freight trains it has to be increased priority of freight trains (expresses). in the overlapping section of rfc 7 and rfc 9 kolín česká třebová there could be generated free capacity due to expresses consolidation (passenger transport), too. acknowledgements this paper is supported by sgs_2016_008. references [1] čech r., šlachtová m. železniční nákladní koridory rfc. železniční magazín czech raildays, m-presse plus, s.r.o. issn 1212-1851, praha, 2015. [2] rfc 7 orient corridor : implementation plan of rail freight corridor 7 orient corridor [online]. [2015-11-20], http://www.rfc7.eu/ckfinder/ userfiles/files/rfc%207%20cid%20all%20books/ rfc%207%20cid%20book%205%20implementation% 20plan%202015-02-16%20approved.pdf. [3] rfc 9 corridor : implementation plan of rail freight corridor 9 [online]. [2015-11-23], http://www. szdc.cz/rfc9/soubory/knihy/05-cid-kniha5.pdf. [4] international union of railways (uic). uic code 406, 2nd edition. paris: uic, 2013. isbn 978-2-7461-2159-1. [5] sždc d1 dopravní a návěstní předpis. praha: sždc, s. o., 321 s., účinnost od 1. 7. 2013. [6] p. šrámek. kapacitní možnosti trati 031 v úseku pardubice hl. n. – hradec králové hl. n. astrophys j pp. 168–178, 1/2015. http: //pernerscontacts.upce.cz/38_2015/sramek.pdf. 62 http://www.rfc7.eu/ckfinder/userfiles/files/rfc%207%20cid%20all%20books/rfc%207%20cid%20book%205%20implementation%20plan%202015-02-16%20approved.pdf http://www.rfc7.eu/ckfinder/userfiles/files/rfc%207%20cid%20all%20books/rfc%207%20cid%20book%205%20implementation%20plan%202015-02-16%20approved.pdf http://www.rfc7.eu/ckfinder/userfiles/files/rfc%207%20cid%20all%20books/rfc%207%20cid%20book%205%20implementation%20plan%202015-02-16%20approved.pdf http://www.rfc7.eu/ckfinder/userfiles/files/rfc%207%20cid%20all%20books/rfc%207%20cid%20book%205%20implementation%20plan%202015-02-16%20approved.pdf http://www.szdc.cz/rfc9/soubory/knihy/05-cid-kniha5.pdf http://www.szdc.cz/rfc9/soubory/knihy/05-cid-kniha5.pdf http://pernerscontacts.upce.cz/38_2015/sramek.pdf http://pernerscontacts.upce.cz/38_2015/sramek.pdf acta polytechnica ctu proceedings 5:59–62, 2016 1 introduction 2 aims 3 materials and methods 3.1 rail freight corridors (rfc) 3.1.1 rfc 7 3.1.2 rfc 9 3.2 the uic 406 leaflet 3.3 the deadline cargo mode 3.4 the simulation method 4 results and discussion 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0012 acta polytechnica ctu proceedings 11:12–15, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app friction buffer stop design petr guziur institute of railway structures and construction, faculty of civil engineering, brno university of technology correspondence: guziur.p@fce.vutbr.cz abstract. friction buffer stops are the favoured construction of buffer stop, mainly due to its high resistance and variety of layout. last but not least is its manner of deceleration induced upon impact and during the braking what makes it smart solution in railway transport safety. the general approach of designing buffer stops is via usage of the kinetic energy and its conversion into work. paper describes input parameters such as train velocity or buffer stop vicinity which is expressed by the safety coefficient implanted within the calculation. furthermore, the paper shows the principle of calculation the friction buffer stop work, or to be more precise, the work of its braking jaws and optionally the work of additional braking jaws located behind the buffer stop. last section of the paper is focused on the examples of designing friction buffer stops, points out the main complications and shows the charts of relation amongst braking distance, kinetic energy and braking force and the charts of relation between deceleration rate and braking distance. keywords: buffer stop, kinetic energy, braking force, braking jaw. 1. introduction with increasing demands on safety in transportation, every threat needs to be considered. this means to keep safe onboard passengers and anyone who can be endangered in case of an accident. this applies even for buffer stops – especially for buffer stops at the dead-end tracks in stations by platforms with pedestrians. the use of fixed buffer stops has become less suitable due to their low resistance and manner of deceleration, which arises in an order of tens of m.s−2 [1] a clearly unacceptable state. according to the czech standard cd z9 [2], it is allowed to design fixed buffer stop (table 1). using those constructions with very limited resistance is justified in particular cases such as deadend tracks in shunting yard, rolling stock depots etc. recently the motivation is to use more reliable constructions to increase the safety in railway transportation. the friction buffer stop with its high resistance and manner of deceleration is more suitable one. buffer stop type 6 cars of 15 t 1 car of 80 t rail buffer stop 1.0 km.h−1 1.6 km.h−1 concrete type "sudop" 0.7 km.h−1 1.1 km.h−1 concrete type "dsb" 1.0 km.h−1 1.6 km.h−1 table 1. resistances of buffer stops used in czech republic [2]. 2. friction buffer stop friction buffer stops have become the preferred constructions at the end of a dead-end track in railway stations abroad (e.g. switzerland, germany [3, 4]). resistance of a friction buffer stop (work of the buffer stop) could be much higher than a fixed buffer stop, and can be customized for various train mass, train speeds and acceptable deceleration rates. disadvantage of a friction buffer stop is obviously its space requirement due to its braking distance (see figure 1). kinetic energy of moving train and its reduction along the braking distance is the criteria to calculate and design the buffer stop work. to be more precise the objective is to design the number of braking jaws and their braking distances. there are two possibilities how to design the friction buffer stop: • friction buffer stop without additional braking jaws (see figure 1); • friction buffer stop with additional braking jaws (see figure 2). figure 1. friction buffer stop [5]. lv – track length required for friction buffer stop lw – braking distance lb – buffer stop length 12 http://dx.doi.org/10.14311/app.2017.11.0012 http://ojs.cvut.cz/ojs/index.php/app vol. 11/2017 friction buffer stop design 3. buffer stop design 3.1. train kinetic energy defining the precise value of kinetic energy of moving rolling stock is to include rotating parts into calculation. to simplify the calculation of train kinetic energy one can use formula 1 for calculation the kinetic energy of non-rotating object of certain mass and velocity. ekin = 1 2 mv2 (1) ekin – train kinetic energy [j] m – train mass [kg] v – train velocity by impact [m.s−1] 3.2. collision speed collision speed is probably the most questionable parameter entering the calculation. in tracks with train control system, which can interfere with driver inappropriate driving (e.g. etcs) – in this case an insufficient braking at the end of dead end track, it is possible to define the collision speed. nevertheless, there are numerous dead-end tracks without any speed controlling system and therefore the driver is responsible for proper braking. collision speeds that could be found in literature are v = 10 km.h−1 or v = 15 km.h−1 (depends on train type or track type) [4–6]. 3.3. safety coefficient for including an inaccuracy entering the calculation of kinetic energy and to ensure the reliability of the design, safety coefficient k has to be counted in calculation. the value of the coefficient varies and depends on type of train, location of dead end track etc. various infrastructure managers approach towards the assessment of the safety coefficient differently. values of safety coefficient usually varies between 1,0 – 2,0 [3, 4, 6]. type of train k [-] freight trains and shunting 1.2 passenger trains, freight trains and shunting, when it is necessary to protect various systems which are located behind or nearby buffer stop 1.5 all trains, where there are traffic zones, structures or residential buildings located behind or nearby buffer stop 1.8 preventing the fall of any train or rolling stock into an abyss 2.0 table 2. safety coefficient [6]. 3.4. buffer stop work in general, the resistance of buffer stop is described as its energy absorbing capacity, therefore the buffer stop work required to stop the train. in case of friction buffer stop, it is the work of braking jaws w with braking forces fb,i,j on braking distances li,j. work of the braking jaws can be obtained using formula 2 and 3. wi = ni ∗ n∑ j=1 fb,i,j ∗ li,j (2) w = s∑ i=1 wi (3) wi – work of i-group of braking jaws [j] ni – number of braking jaws in i-group [-] fb,i,j – braking force of a single braking jaw, based on a braking distance [n] li,j – length of braking distance of i-group of braking jaws, based on lw [m] braking force of each braking jaw decreases along the braking distance [4]. length of braking distance li,j [m] braking force fb,i,j [kn] 0 – 5 40 5 – 8 36 8 – 12 32 12 – 20 28 table 3. braking force of a braking jaw along the braking distance [4]. the design of friction buffer stop (number of braking jaws and their braking distances) is correct if the buffer stop work is equal or higher than the train kinetic energy increased by the safety coefficient: w ≥ ekin ∗ k (4) 3.5. deceleration rate deceleration rate could be the limiting parameter when designing number of braking jaws and their braking distances. the most challenging design comes with dead-end track where parameters of trains vary (mostly the train mass). deceleration rate has to be monitored both upon impact and along the braking distance where additional braking jaws could be designed. fortunately, restriction in deceleration rate applies only to the passenger trains. maximal deceleration rate is set as a = 2.5 m.s−2 [7]. deceleration rate is calculated as follows: amax = fmax m (5) 13 petr guziur acta polytechnica ctu proceedings figure 2. buffer stop with additional braking jaws. amax – maximal deceleration rate [m.s−1] fmax – maximal braking force along the braking distance [n] m – train mass [kg] 3.6. design procedure the design of buffer stop is customized for two types of trains, the lightest one and the heaviest one running on the dead-end track. this approach covers both deceleration rate restriction (usually the limiting factor for light trains) and the highest value of kinetic energy of heavy train (sufficient work of the buffer stop). braking distance for braking jaws on buffer stop is considered as the main braking distance. the tricky part comes with additional jaws behind the buffer stop thus in their braking distances where each pair of additional braking jaws has different braking distance. therefore it is necessary to consider the variable braking force for each pair of additional braking jaws along the braking distance according to table 3. for better understanding, see following examples. 3.6.1. example 1 as it was described in section 3.6, two types of trains were selected for calculation. a light train with mass of 170 t and a heavy train with mass of 300 t. collision speed of trains v = 15 km.h−1, is the same for both of them, and safety coefficient used to increase the kinetic energy k = 1, 8 (see section 3.3). in this situation, the main braking distance is set to 11,3 m (lw – buffer stop) with five pairs of braking jaws and one pair of additional braking jaws with braking distance of 4,0 m (lw – 1. row of additional braking jaws). under those circumstances, the mass of the light train is high enough to design the large number of braking jaws on buffer stop and the limit of deceleration rate is not exceeded. nevertheless, the offset for the first row of additional braking jaws behind the buffer stop must be set. if the light train reaches the additional braking jaws, braking force will increase and therefore the deceleration rate will increase over the limit. for better understanding, see figure 3 and figure 4. figure 3. kinetic energy and braking force along the braking distance (example 1). figure 4. deceleration rate along the braking distance (example 1). 3.6.2. example 2 in this example, the light train with mass of 80 t and the heavy one with mass of 300 t were selected to point the difficulty and variety of design. collision speed of trains v = 15 km.h−1 is the same for both of them and the safety coefficient used to increase the kinetic energy k = 1, 8 was chosen (see section 3.3). in this case, the main braking distance is set to 15,5 m (lw – buffer stop) with just 2 pairs of braking jaws and three rows of additional braking jaws. each row of additional braking jaws consists of 2 pairs of braking jaws. braking distances are designed as follows: 7,0 m (lw – 1. row of additional braking jaws), 5,5 m (lw – 2. row of additional braking jaws) and 4,0 m (lw – 3. row of additional braking jaws). the mass of light train is significantly lower and one has to design fewer braking jaws on buffer stop in order not to exceed the limit of deceleration rate. the offset of first row of additional braking jaws has to be set similarly to example one. 14 vol. 11/2017 friction buffer stop design figure 5. kinetic energy and braking force along the braking distance (example 2). figure 6. deceleration rate along the braking distance (example 2). 4. conclusion the usage of the fixed buffer stops in the railway stations is slowly becoming an anachronism. one can still come across many of those constructions in present railway stations but the latest progression, is to design buffer stops with high resistance and high level of safety. friction buffer stops meet those criteria in every point. the possibilities variety of designing the friction buffer stop is the huge advantage. the range of options varies form light trains with low collision speed, up to heavy trains with high collision speed. the paper discusses the parameters of the buffer stop and shows some examples of design for certain situations. as described above, design buffer stop for some combinations of trains on dead-end track could be challenging while all criteria, such as buffer stop kinetic energy absorbing capacity, restriction of deceleration rate and space options, must be met. acknowledgements the paper was created with support of the project no. lo1408 "admas up – advanced materials, structures and technologies" supported by the ministry of education, youth and sports within the targeted support of program "national program for sustainability i" and student’s grant tender of the faculty of civil engineering, brno university of technology (fast-j-16-3776). references [1] m. bugarín, j.-m. díaz-de-villageas. conception of buffer stop blocks [online]. 2001. [cit. 2016-11-10] "http://www.railway-research.org/img/pdf/329.pdf". [2] c. z9. železniční spodek, vzorový list železničního spodku, zarážedla. české dráhy, s.o., divize dopravní cesty, o.z., praha, 2001. in effect from 1.4.2002. [3] ausführungsbestimmungen zur eisenbahnverordnung / ab-ebv, 2014. artikel: gleisabschluss, bern. [4] db:ds 800 01. bahnanlagen entwerfen – allgemeine entwurfsrichtlinien. [5] i. r. ltd. railway buffer stops planning guidlines [online]. 2009. [cit. 2016-11-05] "http://www.iroads.co.il/sites/default/files/imce/ ir_buffer_stops_guidelines.doc". [6] rve 05.05.31 oberbau, oberbaukonstruktionen, gleisabschlüsse. wien. 2014. [7] guide for the application of tsi loc&pas. version 2.0, 2015. european railway agency, france. 15 acta polytechnica ctu proceedings 11:12–15, 2017 1 introduction 2 friction buffer stop 3 buffer stop design 3.1 train kinetic energy 3.2 collision speed 3.3 safety coefficient 3.4 buffer stop work 3.5 deceleration rate 3.6 design procedure 3.6.1 example 1 3.6.2 example 2 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0135 acta polytechnica ctu proceedings 13:135–141, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app optimization-based approach to tiling of finite areas with arbitrary sets of wang tiles marek tyburec∗, jan zeman czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: marek.tyburec@fsv.cvut.cz abstract. wang tiles proved to be a convenient tool for the design of aperiodic tilings in computer graphics and in materials engineering. while there are several algorithms for generation of finite-sized tilings, they exploit the specific structure of individual tile sets, which prevents their general usage. in this contribution, we reformulate the np-complete tiling generation problem as a binary linear program, together with its linear and semidefinite relaxations suitable for the branch and bound method. finally, we assess the performance of the established formulations on generations of several aperiodic tilings reported in the literature, and conclude that the linear relaxation is better suited for the problem. keywords: wang tiles, binary and continuous linear programming, semidefinite programming, aperiodic tilings, relaxation. 1. introduction wang tiles, squares with colored edges, were invented by and named in honor of hao wang, originally serving as a tool for studying the ∀∃∀ decidability problem of the predicate calculus [1]. wang showed that the decidability problem is equivalent to the domino problem: assume a set of non-rotatable unit-sized (wang) tiles with edges colored according to the ∀∃∀ problem. if it is feasible to cover the infinite plane by translating copies of the tiles, such that all their vertices lie at the integer lattice points of the plane and the adjoining edges share the same color, the problem is said to be solvable; unsolvable otherwise. for an illustration of a sample 2 × 3 wang tiling, see fig. 1. in [2], wang made a fundamental conjecture stating that the domino problem is solvable if and only if the tiling was periodic, i.e., if there existed a rectangular region of the tiling with identically colored horizontal, and vertical edges, respectively. a year later, kahr reduced the turing machine halting problem [3, 4] into the origin-constrained domino problem [5], which implies the domino problem is also undecidable. this can be illustrated by introducing a turing machine for each tile set, halting only if the domino problem is unsolvable. however, there is an infinite number of such tile sets, making the domino problem undecidable as well, and forbidding existence of a general finite algorithm for the generation of infinite tilings. 1.1. aperiodic tile sets undecidability of the domino problem was also proved by wang’s student berger, who used the principle of expanding squares for developing the sets of 20, 426 [6] and 104 tiles [7] that admit only aperiodic tilings of the infinite plane, contrary to the wang fundamental conjecture. ✓ ✗ figure 1. periodic tiling composed of 2 wang tiles over 2 colors. aiming to simplify the berger’s proof, smaller aperiodic sets of wang tiles have been constructed. in 1966, läuchli sent to professor wang an aperiodic set of 40 tiles over 16 colors, but it remained unpublished until 1975 [8]. meanwhile, unaware of läuchli’s result, knuth simplified berger’s set to 92 tiles over 26 colors [9]; and robinson developed sets of 104 and 52 tiles in 1967 [10], of 56 tiles over 12 colors in 1971 [11], and marked existence of a set of 35 tiles [11]. in 1973, penrose developed a new approach based on kites and darts tiling, leading to a set of 34 tiles. robinson, being in contact with penrose, modified the penrose’s approach to reach a reduced set of 32 tiles over 16 colors [12]. using the same technique together with penrose rhombs tiling, grünbaum obtained a set of 24 tiles over 9 colors in 1987 [12]. another two tile sets were discovered due to ammann. in 1978, he used ammann bars to reach 16 tiles over 6 colors [13]. building on the ammann’s a2 tiling, see, e.g., [12], robinson obtained a set of 24 tiles over 24 colors in 1977. subsequent size reduction of the smallest aperiodic set occurred in 1996, as kari developed a new method based on mealy machines multiplying beatty sequences, and presented a set of 14 tiles over 6 colors [14]. čulík, using the same approach, reduced the set even further to 13 tiles over 5 colors [15]. 135 http://dx.doi.org/10.14311/app.2017.13.0135 http://ojs.cvut.cz/ojs/index.php/app marek tyburec, jan zeman acta polytechnica ctu proceedings the search for the minimal aperiodic set is concluded by jeandel, who used a brute-force enumeration approach to find aperiodic sets of 11 tiles over 4 and 5 colors; and proved both that there does not exist an aperiodic set with 10 or fewer tiles and with less than 4 colors [16]. in addition to the classical wang tiles, in 2006, lagae introduced a subset of the wang tiles, the corner tiles, with the connectivity information stored in the colored corners instead of the edges [17]. in the same year, he constructed an aperiodic set of 44 corner tiles over 6 colors [18]. the set was further simplified by nurmi to 30 corner tiles over 6 colors [19]. note that because the corner tiles can be straightforwardly converted to the edge-based wang tiles [17], our approach naturally extends to the corner tiles, see section 4 for a specific example. 1.2. selected applications due to the ability of some tile sets to generate aperiodic patterns, wang tiles received a broad interest among disciplines. also, their original significance for automated theorem proving [2] was supplemented by proofs in cellular automata theory [20]. in 2003, cohen et al. introduced wang tiles to the computer graphics community [21]. since then, wang tiles have been successfully used, e.g., for efficient synthesis of stochastic textures or for generation of poisson disk distributions. molecular realization of wang tiles is due to winfree, who introduced a self-assembly of biological nanostructures into aperiodic crystals [22]. novák was probably the first one who used wang tiles in materials engineering for efficient compression of microstructures [23]; doškář for their reconstruction [24]. doškář further generated large stochastic samples of the alporas® foam and its finiteelement representations in [25]. in [26], tyburec used wang tiles to describe modular assembly of structures, whereas both the topology and arrangement of modules were subjects of optimization. 1.3. tiling generation algorithms existing tiling generation algorithms are designed specifically to the tile sets they handle. in computer graphics, for example, it is essential to generate visually appealing patterns quickly, which is best achieved with stochastic tile sets. tiling a finite area has then a o(n) complexity, adopting, e.g., the stochastic tiling [21], or the hash-based direct stochastic tiling [27] algorithms. the latter algorithm allows straightforward definition of boundary conditions. in the case of mathematical logic, it is crucial to prove that an investigated tile set can tile the whole infinite plane, and does so only aperiodically. as the problem is undecidable, no general algorithm exists, and thus all the algorithms need to exploit the specific structure of each individual (family of) tile sets. although the algorithms are fast, they remain tile set dependent. all tilings of finite-sized areas are restricted to be subsets of the infinite plane, which is not required in the finite domain, and it is almost impossible to introduce boundary conditions. on top of that, there provably exist aperiodic tile sets that admit only non-recursive tiling [28, 29], forbidding design of any problem-specific algorithm to tile the infinite plane. to the authors’ knowledge, the only method that has been used for tiling of finite-sized areas by arbitrary tile sets is the backtracking algorithm, see [30]. although the algorithm is general, it is generally inefficient, as it commonly creates impossible assemblies too early. moreover, distant boundary conditions make the problem more difficult to solve. 1.4. contributions this paper aims to overcome the shortcomings of the methods outlined in the previous section, and to develop an approach that handles tiling of finitesized areas using arbitrary tile sets, together with a straightforward approach to define edgeor tilebased boundary conditions. our exposition is structured as follows. in section 2, we develop a binary linear programming representation of the tiling generation problem. the formulation is relaxed, and its linear and semidefinite approximations are presented. in order to show generality of the formulations, we introduce their extensions to solve the tile-packing problem, to prove that a tile set can not tile the plane, and to prove that a tile set admits periodic tiling. the relaxations are finally employed in a branch and bound method in section 3 and their performance is assessed in section 4. 2. optimization problem formulation 2.1. valid tiling consider a finite rectangular area a of the size nt,h × nt,w, with nt,h and nt,w denoting its height and width, respectively. placing wang tiles (of a unit size) from the tile set t , such that their vertices lie at the integer lattice points of a, we obtain a tiling with nt,h rows and nt,w columns of tiles. each tile k ∈ {1, .. , nt} from the tile set t , with nt denoting the number of tiles, is described by the 4-tuple (nk, wk, sk, ek), assigning color codes c to the north, west, south, and east edge of the tile, respectively. in any valid tiling, for any two adjacent tiles k and `, with the tile ` being placed in the east of k, it holds that ek = w`. (1) similarly, for any two vertically adjacent tiles k and m, with the tile m being placed in the south of k, it stands that sk = nm. (2) both the cases are illustrated in fig. 2. 136 vol. 13/2017 optimization approach to tiling of finite areas with wang tiles sm nm wm em s` n` w` e` sk nk wk ek figure 2. horizontal, and vertical connectivity among tiles k-`, and k-m, respectively. 2.2. binary linear programming formulation let x ∈ {0, 1}nt,h·nt,w·nt denote a binary vector describing the placement of individual tiles within a. the vector consists of all the xi,j,k variables, where i ∈{1, .. , nt,h} and j ∈{1, .. , nt,w} are the row and column iterators, respectively. let us also define xi,j,k = { 0 if the tile k is not at (i, j), 1 if the tile k is at (i, j). (3) because each position is occupied by a single tile, we have ∑ k∈t xi,j,k = 1, ∀i ∈h,∀j ∈w. (4) the compatibility constraint (1) can be written in terms of x after realizing that the number of tiles at (i, j) that have east edge colored by c ∈ c has to be equal to the number of tiles at (i, j + 1) with west edge colored also by c. consequently, we have∑ k∈t xi,j,k[ek = c] − ∑ k∈t xi,j+1,k[wk = c] = 0. (5) as the sums in eq. (5) are either equal to 0 or 1, resulting from eq. (4), only two options are possible: if the edge is colored by c, the relation simplifies to 1 − 1 = 0, (6) otherwise, it equals to 0 − 0 = 0, (7) showing that (5) remains valid in both cases. after applying the same approach to eq. (2) and writing constraints through the entire area a, we obtain the following binary linear program: min x 0 (8a) s.t. ∑ k∈t xi,j,k[ek = c] − ∑ k∈t xi,j+1,k[wk = c] = 0, ∀c ∈c,∀i ∈h,∀j ∈w \{nt,w}, (8b) ∑ k∈t xi,j,k[sk = c] − ∑ k∈t xi+1,j,k[nk = c] = 0, ∀c ∈c,∀j ∈w,∀i ∈h\{nt,h}, (8c) ∑ k∈t xi,j,k = 1, ∀i ∈h,∀j ∈w, (8d) xi,j,k ∈{0, 1}, ∀i ∈h,∀j ∈w,∀k ∈t . (8e) despite the program (8) being formulated straightforwardly, it is hard to solve in this original form due to the (non-convex) integrality constraint (8e). note that the problem is np-complete because of its combinatorial nature [31]. 2.3. linear programming relaxation in order to make the problem (8) easier to solve, let us relax the integrality constraint (8e) into a linear form 0 ≤ xi,j,k ≤ 1, ∀i ∈h,∀j ∈w,∀k ∈t . (9) clearly, eq. (9) allows for all configurations of (8e), but additionally also intermediate non-binary values. in addition, because the constraints in (8) are linear, the resulting approximation is convex and reads as min x 0 (10a) s.t. ∑ k∈t xi,j,k[ek = c] − ∑ k∈t xi,j+1,k[wk = c] = 0, ∀c ∈c,∀i ∈h,∀j ∈w \{nt,w}, (10b) ∑ k∈t xi,j,k[sk = c] − ∑ k∈t xi+1,j,k[nk = c] = 0, ∀c ∈c,∀j ∈w,∀i ∈h\{nt,h}, (10c) ∑ k∈t xi,j,k = 1, ∀i ∈h,∀j ∈w, (10d) 0 ≤ xi,j,k ≤ 1, ∀i ∈h,∀j ∈w,∀k ∈t . (10e) 2.4. semidefinite programming relaxation the integrality constraint (8e) can be rewritten using the non-convex quadratic constraint x2i,j,k − xi,j,k = 0, ∀i ∈h,∀j ∈w,∀k ∈t , (11) which is satisfied only if xi,j,k = {0, 1}, being thus equivalent to (8e). this quadratic form does not simplify the solution, because the problem is still np-complete, but we will use it for the derivation of a semidefinite programming relaxation. let us now substitute the binary variables x ∈{0, 1} with y ∈{−1, 1}, which requires x = 1 2 (y + 1). (12) inserting this into eq. (3) yields yi,j,k = { −1 if the tile k is not at (i, j), 1 if the tile k is at (i, j). (13) consequently, we can write a non-convex quadratically-constrained optimization program in terms of y 137 marek tyburec, jan zeman acta polytechnica ctu proceedings min y 0 (14a) s.t. ∑ k∈t (yi,j,k + 1)[ek = c] − ∑ k∈t (yi,j+1,k + 1)[wk = c] = 0, ∀c ∈c,∀i ∈h,∀j ∈w \{nt,w}, (14b) ∑ k∈t (yi,j,k + 1)[sk = c] − ∑ k∈t (yi+1,j,k + 1)[nk = c] = 0, ∀c ∈c,∀j ∈w,∀i ∈h\{nt,h}, (14c) ∑ k∈t yi,j,k = 2 − nt, ∀i ∈h,∀j ∈w, (14d) y2i,j,k = 1, ∀i ∈h,∀j ∈w,∀k ∈t . (14e) let us further introduce a symmetric matrix y ∈ snt,y·nt,x·nt defined as y = yyt. (15) the definition directly implies that any solution to the quadratically-constrained optimization problem (14) renders the matrix y positive semidefinite, in our notation y � 0, and the rank of the matrix y equal to 1. moreover, the definition of y constrains all the elements in the main diagonal of y equal to 1. consequently, another equivalent formulation to the problem (8) reads as min y,y 0 (16a) s.t. ∑ k∈t (yi,j,k + 1)[ek = c] − ∑ k∈t (yi,j+1,k + 1)[wk = c] = 0, ∀c ∈c,∀i ∈h,∀j ∈w \{nt,w}, (16b) ∑ k∈t (yi,j,k + 1)[sk = c] − ∑ k∈t (yi+1,j,k + 1)[nk = c] = 0, ∀c ∈c,∀j ∈w,∀i ∈h\{nt,h}, (16c) ∑ k∈t yi,j,k = 2 − nt, ∀i ∈h,∀j ∈w, (16d) diag(y) = 1, (16e) y − yyt = 0. (16f) the only non-convex constraint (16f) can be relaxed, see, e.g., [32], into a convex form y − yyt � 0, (17) and based on the schur complement lemma equivalently rewritten to ( 1 yt y y ) � 0. (18) finally, the semidefinite programming relaxation of the binary linear program (8) reads as min y,y 0 (19a) s.t. ∑ k∈t (yi,j,k + 1)[ek = c] − ∑ k∈t (yi,j+1,k + 1)[wk = c] = 0, ∀c ∈c,∀i ∈h,∀j ∈w \{nt,w}, (19b) ∑ k∈t (yi,j,k + 1)[sk = c] − ∑ k∈t (yi+1,j,k + 1)[nk = c] = 0, ∀c ∈c,∀j ∈w,∀i ∈h\{nt,h}, (19c) ∑ k∈t yi,j,k = 2 − nt, ∀i ∈h,∀j ∈w, (19d) diag(y) = 1, (19e)( 1 yt y y ) � 0, (19f) − 1 ≤ y ≤ 1. (19g) 2.5. extensions introduced formulations can include additional requirements for generated tilings. we list some of them below, but only in terms of x, as substitution of eq. (12) into developed equations is obvious. 2.5.1. tile-based boundary conditions there are four types of tile-based boundary conditions. first, we can enforce placement of the tile k ∈t at position (i, j): xi,j,k = 1, i ∈h, j ∈w, k ∈t . (20) conversely, avoiding the tile k at (i, j) requires xi,j,k = 0, i ∈h, j ∈w, k ∈t . (21) the requirement that the same tile is placed at both (i, j) and (p, q) can be written as xi,j,k − xp,q,k = 0, {i, p}∈h,{j, q}∈w,∀k ∈t . (22) enforcing different tiles at (i, j) and (p, q) requires xi,j,k+xp,q,k ≤ 1, {i, p}∈h,{j, q}∈w,∀k ∈t . (23) 2.5.2. edge-based boundary conditions starting from eq. (5), we can also easily formulate edge-based boundary conditions. to constrain the north edge at (i, j) to the color c ∈c, we write∑ k∈t xi,j,k[nk = c] = 1, i ∈h, j ∈w, c ∈c. (24) if the north edge at (i, j) has to differ from c, it holds that∑ k∈t xi,j,k[nk = c] = 0, i ∈h, j ∈w, c ∈c. (25) 138 vol. 13/2017 optimization approach to tiling of finite areas with wang tiles equal color c of two edges, e.g., of the north edge at (i, j) and of the west edge at (p, q), is provided by∑ k∈t xi,j,k[nk = c] − ∑ k∈t xp,q,k[wk = c] = 0, {i, p}∈h,{j, q}∈w,∀c ∈c. (26) finally, different coloring of the north edge at (i, j) and the west edge at (p, q) is guaranteed through∑ k∈t xi,j,k[nk = c] + ∑ k∈t xp,q,k[wk = c] ≤ 1, {i, p}∈h,{j, q}∈w,∀c ∈c. (27) 2.5.3. periodic tiling periodic tiling is a tiling with equally colored both horizontal and also both vertical edges. thus, periodic tiling secures tilability of the infinite plane. building on the edge-based boundary conditions, a periodic tiling can be enforced by∑ k∈t x1,j,k[nk = c] − ∑ k∈t xnt,h,j,k[sk = c] = 0, ∀j ∈w,∀c ∈c, (28a) ∑ k∈t xi,1,k[wk = c] − ∑ k∈t xi,nt,w ,k[ek = c] = 0, ∀i ∈h,∀c ∈c. (28b) 2.5.4. tile packing problem the tile packing problem, see, e.g., [30], consists in generation of a tiling with periodic boundary conditions, and each tile has to be placed exactly once. in order to describe the problem in our framework, we just need to use eq. (28) together with∑ i∈h ∑ j∈w xi,j,k = 1, ∀k ∈t . (29) 2.5.5. objective function obviously, the developed formulations can contain arbitrary convex objective function. for instance, we can minimize the maximum occurrences of tiles, compose the tiling such that it fits some pattern, etc. 3. the branch and bound method the whole integer design space, i.e., {0, 1} for the linear (10) and {−1, 1} for the semidefinite (19) approximation, can be described using a tree data structure, with each node corresponding to variables x, or y. however, as there is exactly one tile per position, it is advantageous to use the nt-ary tree representation instead of the binary one, so that the nodes correspond to the positions and branch into nt children, avoiding solution to infeasible programs placing multiple tiles at the same position. solution to the developed approximations, (10) or (19), corresponds to exploring the root node of the tree. unfortunately, because the approximations widen the feasible design space, their solution does not assure to solve the original problem (8) due to the design variables being allowed to take non-integer values. in order to overcome that, we branch some nodes of the tree, i.e., gradually fix all variables that correspond to the specific tile position to all possible combinations of integer values, nt in our case. the relaxation is solved in each node, bounding the problem1, hence the name of the method branch and bound [33]. the branching and bounding continues until the optimal (feasible) solution to (8) is found. if no such solution exists, the solution space is proven to be empty and the tiling generation problem infeasible. without boundary conditions, infeasibility proves the tile set does not tile infinite plane. 3.1. branching rule in our implementation, we firstly branch the most promising nodes, in which the integer infeasibility of the convex relaxation is minimal. for the linear relaxation, we have iinf = ∑ i∈h ∑ j∈w ∑ k∈t (xi,j,k − x2i,j,k), (30) or iinf = nt,hnt,wnt − ∑ i∈h ∑ j∈w ∑ k∈t y2i,j,k (31) for the semidefinite relaxation. if iinf = 0, a feasible solution to the original problem was found. 3.2. variables ordering in each node to be branched, we have to define variables that will be fixed. it seems to be advantageous to fix the variables corresponding to the most difficult position, i.e., to the position with the highest integer infeasibility. in the case of the linear relaxation, the to-be-fixed position (i, j) requires max ∀i∈h,∀j∈w ∑ k∈t (xi,j,k − x2i,j,k). (32) similarly, for the semidefinite relaxation we write max ∀i∈h,∀j∈w (nt − ∑ k∈t y2i,j,k). (33) 4. examples building upon the previous sections, we implemented a custom branch and bound algorithm in c++. the linear relaxations are solved using gurobi [34], the semidefinite programming relaxations are optimized by mosek [35]. as gurobi contains a build-in state-ofthe-art branch and cut algorithm, usable only for the linear relaxations, we also measured its performance. five tile sets were tested altogether: jeandel’s 11 tiles over 4 colors [16], čulík’s 13 tiles [15], ammann’s 16 tiles [12], lauchli’s 40 tiles [12], and nurmi’s 1bounding occurs only if an objective fuction is employed. 139 marek tyburec, jan zeman acta polytechnica ctu proceedings 11(4) [16] 13(5) [15] 16(6) [12] 40(16) [8] 30(6) [19] tile set tiling sample lp 5 × 5 0.13 s/12 rel. 0.10 s/14 rel. 0.10 s/17 rel. 0.01 s/1 rel. 0.01 s/1 rel. sdp 5 × 5 12.65 s/78 rel. 11.07 s/27 rel. 21.66 s/33 rel. 360.40 s/121 rel. 126.17 s/31 rel. lp 6 × 6 0.10 s/12 rel. 0.12 s/14 rel. 0.13 s/17 rel. 0.02 s/1 rel. 0.03 s/1 rel. sdp 6 × 6 69.60 s/34 rel. 227.24 s/79 rel. 184.23 s/33 rel. 1517.59 s/121 rel. 1522.72 s/91 rel. lp 15 × 15 151.82 s/2190 rel. over 6000 s 2.15 s/17 rel. 8.52 s/41 rel. 0.13 s/1 rel. lpg 15 × 15 0.06 s/0 rel. 0.09 s/0 rel. 0.39 s/0 rel. 0.79 s/0 rel. 1.49 s/0 rel. lp 20 × 20 1545.94 s/26137 rel. over 6000 s 8.52 s/17 rel. 32.09 s/41 rel. 4.23 s/31 rel. lpg 20 × 20 0.08 s/0 rel. 0.12 s/0 rel. 0.79 s/0 rel. 1.57 s/0 rel. 8.01 s/0 rel. lpg 25 × 25 396.5 s/7368 rel. 0.12 s/0 rel. 1.34 s/0 rel. 2.96 s/0 rel. 23.95 s/0 rel. lpg 30 × 30 390.18 s/3821 rel. 1361.51 s/5950 rel. 2.31 s/0 rel. 5.23 s/0 rel. 48.97 s/0 rel. table 1. time demands and solved relaxations count in generation of aperiodic tilings. lp and sdp denote custom developed branch and bound method supplied with the linear and semidefinite relaxation, respectively; lpg stands for gurobi cut-and-branch method and the linear relaxation. set of 30 corner tiles [19]. the tile sets were selected to sample different construction methods, tile set sizes, and numbers of colors used. results of the benchmarks, ran on the intel® core™ i5-4210h processor, are summarized in table 1. they indicate that the linear relaxation surpassed the semidefinite one in terms of speed and in the number of explored nodes. the latter results from the property of the simplex algorithm, used for the solution of linear relaxations, to terminate in vertices of the feasible space polytope. these are more likely integer-feasible than an interior point found by the interior-point algorithm, used for the solution of semidefinite programs. comparison of our branch and bound algorithm and gurobi’s build-in branch and cut, both using the linear relaxation, is even clearer. our implementation lacks heuristics, parallelism, and generation of cuts, consequently being inefficient for tile sets with large number of degrees of freedom, such as the čulík one. 5. conclusions in this contribution, we demonstrated that the tiling generation problem of finite-sized areas can be formulated as an integer optimization problem; and we also introduced its linear and semidefinite programming relaxations. both the relaxations were employed in the branch and bound method and benchmarked on five tile sets. the results indicate that the linear relaxation is more suitable for practical applications. 6. nomenclature list of symbols a rectangular area c color set h set of rows t tile set w set of columns c color iterator ek east edge color of the tile k i row iterator j column iterator k, `, m tile iterators nc number of colors in c nk north edge color of the tile k sk south edge color of the tile k nt,h number of rows in h nt,w number of columns in w xi,j,k design variable, xi,j,k ∈{0, 1} yi,j,k design variable, yi,j,k ∈{−1, 1} wk west edge color of the tile k iinf integer infeasibility 140 vol. 13/2017 optimization approach to tiling of finite areas with wang tiles acknowledgements this work was supported partly by the grant agency of the ctu in prague, sgs17/042/ohk1/1t/11, by the grant agency of the czech republic, gačr 15-07299s, by the ministry of industry and trade, mpo fv10202, and by the technology agency of the czech republic, tačr th02020420. references [1] h. wang. dominoes and the aea case of the decision problem. symposium on the mathematical theory of automata pp. 23–55, 1962. 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http://dx.doi.org/10.1002/9781118627372 http://www.gurobi.com http://docs.mosek.com/8.1/cxxfusion acta polytechnica ctu proceedings 13:135–141, 2017 1 introduction 1.1 aperiodic tile sets 1.2 selected applications 1.3 tiling generation algorithms 1.4 contributions 2 optimization problem formulation 2.1 valid tiling 2.2 binary linear programming formulation 2.3 linear programming relaxation 2.4 semidefinite programming relaxation 2.5 extensions 2.5.1 tile-based boundary conditions 2.5.2 edge-based boundary conditions 2.5.3 periodic tiling 2.5.4 tile packing problem 2.5.5 objective function 3 the branch and bound method 3.1 branching rule 3.2 variables ordering 4 examples 5 conclusions 6 nomenclature list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0085 acta polytechnica ctu proceedings 13:85–88, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app water absorption capacity coefficient and mass moisture of rammed earth material barbora mužíková∗, tereza otcovská, pavel padevět czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: barbora.muzikova@fsv.cvut.cz abstract. this article presents the development of mass moisture of rammed earth material and determines the water absorption capacity coefficient for rammed earth with illite-kaoline clay. specimens of two prescriptions were rammed in the moulds. they were settled in the box with soft foam that was moistened. the level of moistening was kept constant. the specimens were regularly measured and weighted. two measuerements were carried out – one of montmorillonite clay during 13 days and one of illite-kaoline clay during three hours. the development of moisture increase was captured and the water absorption capacity coefficient for illite-kaoline clay was determined and compared to coefficients of common building materials. keywords: rammed earth, illite-kaoline clay, mass wetness, water absorption capacity coefficient. 1. introduction this article presents earth material as a modern building material based on the long history of using, but in last centuries it was replaced by modern materials. the material characteristics were never fully researched due to the replacement. the demand for the earth material is increasing nowadays following the principles of sustainable buildings development [1, 2]. the main advantages of using earth as a building material are good air humidity balance, lower costs for material and construction, and also the fact that the earth based material has a positive effect on human health. rammed earth is accessible, environmentally friendly and suitable for low-cost building in person. according to the sustainable buildings development not only the mechanical properties are important in the structure design. the quality of building material is also given by economic and socio-cultural criteria. for example the rammed earth primary energy consumption connected with production of building material is about 44 kwh/m3. this is 18× less than for prefabricated concrete material [3–6]. the increasing interest in unburned earth material can be seen in the sales of unburned clay bricks in portfolio of heluz company shown in table 1 [7]. the mechanical properties of the unburned earth depend mainly on the composition and the method of processing. the basic methods of using earth are mudwalls, rammed earth, earth bricks, earth infill in a timber frame construction etc. the material which the research is focusing on is rammed earth. rammed earth walls are constructed by the compacting (ramming) of moistened subsoil into place between temporary framemwork panels. the process of building is shown in figure 1. first of all, the reinforced plywood framework is built and then the first layer moist earth (mixture of sand, gravel year sold unburned bricks [pcs] 2010 700 2011 8,600 2012 19,800 2013 21,500 table 1. number of sold unfired bricks of heluz company [7]. figure 1. the process of building rammed earth wall. and clay) is filled in. secondly, the layer of moist earth is compressed by a pneumatic backfill tamper. thirdly, next layer of moist earth is added and then succesive layers of moist earth are added and compressed. finally, the framework is removed leaving the rammed earth wall with visible layers of compacted earth. when dried, the result is a dense, hard monolithic wall [4–6]. 85 http://dx.doi.org/10.14311/app.2017.13.0085 http://ojs.cvut.cz/ojs/index.php/app b. mužíková, t. otcovská, p. padevět acta polytechnica ctu proceedings figure 2. the production of rammed earth specimens. 2. testing of unfired earth material 2.1. rammed earth and water the rammed earth material is considered for building bearing walls of construction so that the material is often exposed to rain and moisture of external conditions. all materials with open porous structures like loam are able to store and transport water within their capillaries. water always travels from regions of higher humidity to regions of lower humidity. the value of the water absorption coefficient w is known for common building materials, some of them are shown in table 2. in the research the value for rammed earth with illite-kaoline clay was searched for to be compared with values of other building materials. the premise was that the value of the water absorption coefficient w for rammed earth would be higher than for common building materials. 2.2. preparing material prescriptions for the specimens were designed of clay, sand and water. there were two prescriptions gem i and gem ii of montmorillonite clay and agl i and agl ii of illitic-kaolinite clay. prescriptions i and ii differed in the sand/clay ratio. the amount of the components was expressed by the weight percent. the amount of water was settled as a water/clay ratio, same at both prescriptions, all prescriptions are shown in table 3. figure 3. specimens in the climatic chamber. material w ρ [kg/m2 √ hour] [kg/m3] silty loam 3.7 1900 clayer loam 1.6 1940 cement concrete 1.8 2290 hollow brick 8.9 1165 solid brick 25.1 1750 table 2. water absorption capacity coefficient of common building materials [5]. first of all, the sand was well mixed with two thirds of water. secondly, the clay and the rest of water was added. it was mingled by a drilling machine with a special ending. specimens of size 40×40×160 mm were prefabricated for the testing in the laboratory. the production was made by ramming into moulds by hand and by drilling machine. the moulds were wiped by oil and the earth was rammed in four to five layers. layers were compresed by a mechanical tamper and the last one was always compresed by hand. produced specimens are shown in figure 2. specimens were tested after four weeks from production. by this time the humidity from the prefabrication is stable and is not changing any more. the hardening process of clay is based only on the drying out the mixed water. during these four weeks the specimens were settled in constant condition of temperature 20 °c and moisture 50 % in the climatic chamber (specimens in climatic chamber figure 3). 2.3. practical tests firstly, specimens were put in the covering of nylon to avoid the erosion of particles from the submerged surface. secondly, they were weighted and finally settled in the box with soft foam that was moistened. the level of moist was constant, the water was added during the test. the specimens were regularly measured and weighted. some of the specimens were broken during the test, they were removed from the testing and not evaluated. 86 vol. 13/2017 water absorption capacity coefficient of rammed earth material prescription type of used clay sand [%] clay [%] water/clay ratio [-] bulk density [kg/m3] agl i illite-kaoline 80 20 0.37 2191 aglii illite-kaoline 75 25 0.37 2159 gem i montmorillonite 80 20 0.37 2073 gem ii montmorillonite 75 25 0.37 2040 table 3. prescriptions of used specimens and bulk densities of final specimens. twenty specimens of gem and ten specimens of agl were tested. the test of gem was 14 days long and the agl was evaluated for three hours of the test beginning. 3. evaluation of the measurement 3.1. mass moisture development the mass moisture mw was calculated from measured weigh using relation: mw = m−md md . (1) the mass moisture mw was calculeted as a weight of the wet specimen m minus the weight of dry specimen md and then it was divided by the weight of the dry specimen md (the weight of the specimen in the beginning of the test) and it was multiplied by 100 to transform it to percentage of weight. the data of specimen with different starting weight can be compared with the mass moisture. the data represented the specimens that were not broken during the whole testing. they had the same conditions at the same time. the mass moisture increase is shown in the figure 4. the mass moisture in percent is shown on the vertical axis and the time (in hours or days) is shown on the horizontal axis. 3.2. water absorption capacity coefficient the water absorption cofficient w determines the ability of material absorp water by an area in defined time period. the formula for its calculation is: w = w √ t . (2) units of w are kg/m2 √ t. the quantity of water w that can be absorbed by an area a is defined by the formula: w = m−md a . (3) units of weight of the wet specimen m and weight of the dry specimen md are in kg and the time t is measured in hours. the weights were measured in a time period of one hour for agl prescriptions. the absorbing area of specimens is 40×40 mm. the coefficient w was calculated for prescription agl i and agl ii. figure 4. mass moisture development of prescriptions agl and gem. 4. discussion the mass moisture development of agl (blue column) and gem (red column) is shown in figure 4. the mass moisture development of agl i rose to 1.39 ± 0.28 % and of agl ii to 1.32 ± 0.16 % after four hours. the biggest increase was in the first hour of testing then the mass moisture rose quite slower. the specimen of gem i rose to 14.89 ± 0.13 % and of gem ii to 16.70 ± 0.92 % after 14 days. the trend of development was simillar to the first measurement, the biggest change was in the first days then it slowed down. the coefficient w was calculated to 2.80±0.55 kg/m2 √ t for agl i and 2.70±0.30 kg/m2 √ t agl ii. the calculated values were compared to known characteristics of common building materials as it is shown in figure 5. the coefficients w for two types of loam, cement concrete and two types of bricks were compered to 87 b. mužíková, t. otcovská, p. padevět acta polytechnica ctu proceedings figure 5. water absorption capacity coefficient of common building materials and tested rammed earth agl. illitic-kaolinite rammed earth material. the results of made tests agl are between the values of silty and clayey loam. the coefficient w for agl i is only 1.56× higher than for cement concrete. the value of agl i is 0.11× lower in comparation to the solid brick. the value above 10 kg/m2 √ t is consider to be not so good for building material. non of the prescriptions exceeds this limit, so in the point of view of the coefficient w, the rammed earth material with illite-kaoline clay is competitive to other common building material. 5. conclusion the rammed earth material properities in relation to mass moisture and water absorption capacity coefficient was observed. two kinds of specimens were prepared, one prescription of illite-kaoline clay and the other of montmorillonite clay. the increasing of mass moisture in rammed earth of illite-kaoline clay in three hours and montmorillonite clay in forteen days was observed. the progress of mass moisture increasing was fastest in the beginnig of moistenig for both cases. the other observed value was the water absorption capacity coefficient w for rammed earth with illite-kaoline clay (prescriptions agl i and agl ii). the calculated values are comperative to other buiding material as cement concrete and they are according to the values found in literaterature [5] for silty loam and clayey loam. from that point of view, rammed earth have better properities than for example the solid brick. generally the calculated values of water absorption capacity coefficient are lower than it was expecting in the begining of the test. acknowledgements this research was financial supported by the faculty of civil engineering at ctu in prague (sgs project no. 16/201/ohk1/3t/11). authors gratfully aknowledge free supply of material necessary for experimental measurement to lb minerals, s.r.o. company. references [1] p. suske. hlinené domy novej generacie. first printing. alfa bratislava, 1991. [2] p. suske. nepalena hlina v moderni architekture. dokument era21: ekologie, realizace, architektura (2):69–70, 2004. [3] b. little, m. morton. building with earth in scotland: innovative design and sustainability. first printing. scottish executive central research unit, 2001. [4] j. norton. building with earth, a handbook. first printing. intermediate technology development group limited, 1986. [5] g. minke. building with earth design and technology of sustainable architecturer. second printing. birkhauser publishe, 2006. [6] i. zabickova. hlinene stavby. first printing. era vydavatelstvi, 2002. [7] i. zabickova, t. otcovska. compressive strength of unburned clay masonry. applied mechanics and materials pp. 31–34, 2016. 88 acta polytechnica ctu proceedings 13:86–89, 2017 1 introduction 2 testing of unfired earth material 2.1 rammed earth and water 2.2 preparing material 2.3 practical tests 3 evaluation of the measurement 3.1 mass moisture development 3.2 water absorption capacity coefficient 4 discussion 5 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0048 acta polytechnica ctu proceedings 22:48–51, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app analysis of the influence of transducer frequency on the ultrasonic measurement of concrete homogeneity dalibor kocáb∗, petr misák, barbora jindrová, martin alexa, tomáš vymazal brno university of technology, faculty of civil engineering, veveri 331/95, 602 00 brno, czech republic ∗ corresponding author: dalibor.kocab@vutbr.cz abstract. the paper analyses the influence of transducer frequency on the determination of concrete homogeneity using the ultrasonic pulse velocity test. transit time measurements were made on a 590 × 590 mm concrete slab, with 110 mm in thickness, in a raster of 5 × 5 points, which means the slab was tested in 25 places. the tests were made using a pundit pl-200 ultrasonic tester using transducers set at 54, 82, and 150 khz. two types of measurements were performed – spot measurements of the ultrasonic pulse transit time at each point and full area scanning. the paper is concluded by an evaluation of the concrete slab’s homogeneity measured by different transducers and techniques in addition to a statistical analysis of how the results are affected by the transducer frequency. keywords: ultrasonic pulse velocity test, concrete, concrete homogeneity, transducer frequency. 1. introduction the ultrasonic pulse velocity test, alongside the rebound hammer test discussed e.g. in [1], is one of the most widespread testing methods in civil engineering [2]. the principle of the ultrasonic pulse velocity test is sending ultrasonic pulses into the material and measuring their transit time. the pulses travel through different materials at a different velocity, depending on their properties [3]. in concrete it is generally true that the pulse velocity is lower in poor quality concrete or in damaged areas (including micro-defects) [4]. ultrasonic pulse transit time is regulated by čsn en 12504-4 [5] and čsn 73 1371 [6], and can be used for measuring: • concrete homogeneity, • the presence of cracks or air voids, • material properties – modulus of elasticity, compressive or tensile strength, • changes in the above-listed properties over time (e.g. by degradation). some of the main advantages of the ultrasonic pulse velocity test is that it is completely non-destructive, which makes it suited for repeated measurements in the same place and convenient for testing in a laboratory or on site. civil engineering usually uses frequencies of 20 – 150 khz; sometimes even up to 500 khz. higher frequencies improve the instrument’s resolution, making the measurement more accurate, but also substantially weakening the pulse, which asks for a compromise depending on the measured element’s thickness as well as other parameters [7]. the motivation for this experiment was the fact that a vikasonic ultrasonic tester [8] was procured by the co-authors [9]. this device is primarily designed for the non-destructive determination of the setting time in cement pastes or for monitoring the progress of their setting and hardening at an early age, which is something that has been addressed in a number of papers [10, 11]. the vikasonic uses transducers that operate at 54 khz. a specimen has the shape of the vicat ring and can be made of a paste, mortar, or concrete, which makes this device rather suitable. the vicat ring, however, is only 40 mm tall and thus the transducers are positioned at this distance as well. with a measuring base of this size, however, it would be better to use higher transducer frequencies [5, 6]. the goal of the paper was to determine how the transducer frequency affects the ultrasonic pulse transit time in concrete during homogeneity measurement. 2. experiment because it offers a wider range of settings during transit time measurement, the experiment was performed with a pundit pl-200 [12] instead of a vikasonic, since they both operate on the same principle. the actual measurement involves repeated sending of ultrasonic pulses and detecting them as they travel through the material. when applied to concrete, the ultrasonic pulse velocity test usually uses two transducers (a transmitter and a receiver) placed on opposite sides of the specimen or structure, measuring the time it takes for the ultrasonic pulse to travel from one transducer to the other, which is then used to calculate the velocity vl [7]. measurements were performed on a concrete slab with 110 mm in thickness and 590 × 590 mm in width and length using three types of transducers – 54, 82 and 150 khz. a 5 × 5 raster of 25 measurement places was laid out on the slab, see fig. 1. the slab was 48 http://dx.doi.org/10.14311/app.2019.22.0048 http://ojs.cvut.cz/ojs/index.php/app vol. 22/2019 analysis of the influence of transducer frequency. . . figure 1. spot measurement layout. scanned three times using each pair of transducers in an area scanning mode. this is a mode that the pundit pl-200 offers and is able to draw the quality distribution over an area of the concrete being tested. the results take the form of contour lines that show different areas in different colours based on the ultrasonic pulse velocity vl. even during the measurement, one can see which areas are of better or worse quality (in other words, where the ultrasonic pulses travel more quickly or slowly). the area scanning mode was used to determine the velocity vl 75 times – three measurements in the raster of 25 points, using all three transducer types. then the pundit pl-200 was switched to a mode measuring the ultrasonic pulse transit time t0. given the fact that the device does not support 82 khz transducers, the points of the raster on the test slab were measured using only transducers of 54 and 150 khz. otherwise the measurement procedure was the same as during area scanning (the sequence of measuring points) and both transducer frequencies were tested three times. thus, 75 values of transit time t0 were measured, from which 75 values of vl were calculated (three per measuring point). aside from homogeneity testing, a statistical analysis of measurements by each transducer was performed. the homogeneity testing followed čsn 73 2011 [13], where the key criterion is the value of the coefficient of variance vl, which, in homogeneous concrete, must not exceed 4.0%. 3. results and discussion the statistical comparison of test results was made following procedures of the analysis of variance (anova) and a two-sample t-test on a level of significance of 0.05. these statistical tests were set up so as to enable the assessment of the mean values of the test results; in case this hypothesis is rejected, the sets of results can be considered significantly different. first, an assessment was made whether the results are comparable at repeated measurements and unchanged instrument settings. no statistically significant deviation was found there. fig. 2 shows the results in the form of contour lines created from the average value of vl for the corresponding point of measurement. for the purposes of further processing, the results were compiled into five files of 75 values of vl. additionally, data normality was tested on a level of significance of 0.01. histograms of the test results are shown in fig.3. fig. 4 shows a box plot of the values of vl. in the case of spot measurements using 54 and 150 khz transducers, there was a statistically significant deviation in the values of vl, which was also confirmed by one-way hypothesis testing. it can thus be said that using 150 khz transducers produces much higher values of vl than using 54 khz ones. the same conclusion was also reached by an analysis of the results of area scanning, which says that measurements with 150 khz transducers return values that are statistically more significant than measurements made with 82 khz transducers. the values of vl measured by 54 or 82 khz transducers, however, do not show any difference. in addition, data from spot and area scanning at the same frequency were compared. no statistically significant deviation was found there. it can thus be stated that the resulting values of ultrasonic wave velocity is influenced primarily by transducer frequency. however, it must be noted that the experiment outcomes may have been affected by the small measurement base or the non-homogeneity of the slab being tested. this would explain why the coefficient of variance exceeded 4.0% in all measurements. 4. conclusion the experiment invites the following conclusions: • transducer frequency has a statistically significant role in determining the final value of ultrasonic pulse velocity vl in concrete. • the precise value of vl thus cannot be precisely ascertained, nor can it be used to calculate any property of concrete (e.g. dynamic modulus of elasticity), which corresponds with the recommendations in astm c597-16 [14]. • on the other hand, the ultrasonic pulse velocity test appears well-suited for determining concrete homogeneity, since it only observes variability in the values of vl. this method can just as well be used for determining the progress of degradation in concrete, e.g. due to exposure to freeze-thaw cycles, as the absolute value of vl is not deciding; only its decrease over time. • the vikasonic can be used for measuring the setting time of cement composites. however, given the small measurement base and the low frequency of 49 d. kocáb, p. misák, b. jindrová et al. acta polytechnica ctu proceedings figure 2. a contour diagram of the test results of area and spot scanning. figure 3. test result histograms. 50 vol. 22/2019 analysis of the influence of transducer frequency. . . figure 4. box plots of test results. transducers (40 mm and 54 khz), this instrument is not entirely suitable for determining material properties such as the modulus of elasticity. • any ultrasonic device thus appears ill-suited for measuring the elastic modulus. there is a possible option of correcting the values based on some other simultaneous measurement (e.g. loading under compression). acknowledgements this paper has been written as part of the project no. ga17-14302s “experimental analysis of the early-age volume changes in cement-based composites”, supported by the ga0 czech science foundation and project no. fast/fch-j-18-5595 “reduction of chemical and autogenous shrinkage of alkaline activated slag”. references [1] p. cikrle, d. kocáb, b. kucharczyková, o. anton. in situ determination of the elastic modulus of concrete by means of ultrasonic pulse method. in 24th concrete days 2017, vol. 272 of solid state phenomena, pp. 70–75. trans tech publications, 2018. doi:10.4028/www.scientific.net/ssp.272.70. [2] v. m. malhotra, n. carino (eds.). handbook on nondestructive testing of concrete. crc press inc., 2nd edn., 2004. [3] v. özkan, i. h. sarpün, s. tuncel. nondestructive testing of materials and structures, vol. 6 of rilem bookseries, chap. relative effects of porosity and grain size on ultrasonic wave propagation in marbles, pp. 373– 379. springer, 2012. doi:10.1007/978-94-007-0723-8_54. [4] a. m. neville. properties of concrete. pearson, new york, 5th edn., 2011. [5] čsn en 12504-4 testing concrete part 4: determination of ultrasonic pulse velocity. standard, čni, prague, 2005. [6] čsn 73 1371 non-destructive testing of concrete – method of ultrasonic pulse testing of concrete. standard, únmz, prague, 2011. [7] b. kopec, et al. nedestruktivní zkoušení materiálů a konstrukcí: (nauka o materiálu iv). akademické nakladatelství cerm, brno, 2008. [8] s. t. systems. vikasonic measuring the early setting and hardening with ultrasonic waves. http: //www.schleibinger.com/cmsimple/en/?setting_ and_maturity:ultrasonic_setting_measurement. [9] i. rozsypalová, o. karel, b. kucharczyková, et al. monitoring of the setting and early hardening with ultrasonic waves. in rehabilitation and reconstruction of buildings, vol. 776 of key engineering materials, pp. 51–54. trans tech publications, 2018. doi:10.4028/www.scientific.net/kem.776.51. [10] a. gibson, d. ciancio. nondestructive testing of materials and structure, vol. 6, chap. early-age ultrasonic testing of concrete and shotcrete using embedded sensors, pp. 485–490. springer, dordrecht, 2013. doi:10.1007/978-94-007-0723-8_69. [11] n. de belie, c. grosse, j. kurz, h.-w. reinhardt. reinhardt hw ultrasound monitoring of the influence of different accelerating admixtures and cement types for shotcrete on setting and hardening behavior. cement and concrete research 35:2087–2094, 2005. doi:10.1016/j.cemconres.2005.03.011. [12] proceq. pundit ultrasonic pulse velocity and pulse echo tomography. https://www.proceq.com/compare/ pundit-ultrasonic-pulse-velocity-and-pulseecho-testing/. [13] čsn 73 2011 non-destructive testing of concrete structures. standard, únmz, prague, 2012. [14] astm c597-16 standard test method for pulse velocity through concrete. standard, astm international, west conshohocken, 2016. 51 http://dx.doi.org/10.4028/www.scientific.net/ssp.272.70 http://dx.doi.org/10.1007/978-94-007-0723-8_54 http://www.schleibinger.com/cmsimple/en/?setting_and_maturity:ultrasonic_setting_measurement http://www.schleibinger.com/cmsimple/en/?setting_and_maturity:ultrasonic_setting_measurement http://www.schleibinger.com/cmsimple/en/?setting_and_maturity:ultrasonic_setting_measurement http://dx.doi.org/10.4028/www.scientific.net/kem.776.51 http://dx.doi.org/10.1007/978-94-007-0723-8_69 http://dx.doi.org/10.1016/j.cemconres.2005.03.011 https://www.proceq.com/compare/pundit-ultrasonic-pulse-velocity-and-pulse-echo-testing/ https://www.proceq.com/compare/pundit-ultrasonic-pulse-velocity-and-pulse-echo-testing/ https://www.proceq.com/compare/pundit-ultrasonic-pulse-velocity-and-pulse-echo-testing/ acta polytechnica ctu proceedings 22:48–51, 2019 1 introduction 2 experiment 3 results and discussion 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0112 acta polytechnica ctu proceedings 12:112–115, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app dangerous driver’s behavior libor topol∗, ivo drahotský univerzita pardubice, dopravní fakulta jana pernera, studentská 95, 532 10 pardubice ∗ corresponding author: libortopol@seznam.cz abstract. different types of dangerous driver’s behavior are discussed in this paper. case study analysis of young drivers’ hazardous behavior is also provided from authors’ practices. the practical research is collaborated with the statistical data that aggressive driving induces the number of fatalities on the roads in the czech republic by 25%. the improvement to driver’s training and other countermeasures is then suggested. keywords: dangerous driving behavior, driver training, road safety. year number of accidents number of fatalities number of serve injuries number of light injuries 2010 75522 753 2823 21610 2011 75137 707 3092 22519 2012 81404 681 2986 22590 2013 84398 583 2782 22577 2014 85859 571 2409 18074 2015 93067 660 2540 24426 table 1. evolution of basic indicators of accidents in 2010-2015. 1. introduction road transport occupies a significant segment in transportation in general and due to its high accident rates and hence economical losses, considering the affecting factors is rather crucial for safety improvement. road traffic safety may be characterized by direct and indirect safety indicators. direct indicators are the number of road accidents, their consequences and economical losses. indirect indicators are phenomena or circumstances that project safety assessment into experimental plane of relations between human behavior and transportation safety (following the speed limits, application of safety systems, maintaining safe distances between vehicles, estimation and anticipation of errors in real traffic). high traffic intensity on roads is a reason of high workload for car drivers. this overload is related to the number of accidents in the czech republic, which, despite various measures for its reduction in comparison with other european countries, is still very high. traffic accidents are a major social issue. the table tracks a noticeable increase of fatalities in traffic accidents. number of light and heavy injuries increases as well. one may notice a constant increase in total number of accidents. in 2015 total of 93,067 accidents was figure 1. total number of recorded accidents in the czech republic from 2011-2015. reported and investigated. for example, compared to 2014, there were about 7,208 more accidents in 2015. every day the czech police has conduct around 255 accident investigations on average. this is equal to approximately 14.901.712,00 czk of property losses every day[1]. 2. dangerous driver’s behavior among other factors, aggressive driving contributes to traffic accidents to a high extent. per úamk, 91% of drivers have been exposed to aggressive driving to a certain extent. under the term of dangerous driver’s behavior, we may assume the one that is in contradiction to official and so-called unofficial or nonwritten rules, an unexpected, aggressive or unusual and non-customary behavior on the road[2]. human behavior is diverse. driver’s behavior differs from every person. it is possible to categorize dangerous behavior from different perspectives. inattentive driving can be related to dangerous behavior as it is, besides speeding, the most common cause of traffic accidents. another category is driving while being involved in some secondary activities that are not related to vehicle control. these are smoking, talking 112 http://dx.doi.org/10.14311/app.2017.12.0112 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 dangerous driver’s behavior accident #1 accident #2 time of accident: 5:15 am time of accident: 6:54 road class: ii road class: ii road section type: roundabout road section type: two-way road weather conditions 8°c, clear weather conditions 12°c, clear driver’s age: 25 y.o. driver’s age: 28 y.o. vehicle vehicle car make: audi a6 car make: suzuki swift year of issue: 2001 year of issue: 2004 outcomes outcomes # of fatalities: 3 – immediate (including driver), 1 – in hospital # of fatalities: 3 – immediate (including driver) # of severely injured: 2 # of severely injured: 3 description: driver approached and entered the roundabout without any or with low speed reduction and thus couldn’t manage the turning but ran across the safety island in the middle of roundabout and broke through the roundabout fence ended up on a roof in retention tank. accident sketch can be seen in the image bellow. description: driver exit the lane and with right vehicle side hit front of vehicle in an opposing traffic. accident sketch can be seen in the image bellow. table 2. figure 2. damaged vehicles from accident 1 (left) and accident 2 (right). 113 libor topol, ivo drahotský acta polytechnica ctu proceedings on the phone while driving, eating, drinking, involving into conversations with other passengers and so on. a driver’s personality, psychological and medical condition, lack of knowledge of traffic rules can be the risk itself[3]. these and many features of driver’s behavior that can be gathered under one term of aggressive driving. committing traffic violations can occur both because of negligence, and unintentionally[4]. driver’s age is also one of the factors that influences road safety. lack of experience of novice drivers can cause a potential danger in traffic situations as well as deteriorating physical state of senior drivers[5]. hence, we can categorize dangerous behavior by age or driving experience. young drivers (age group under 24 years old) are said to be the highest risk group of drivers in agewise[6]. they lack the capacity to adequately evaluate dangerous situations, they tend to overestimate their own abilities, have little driving experience and, therefore, fewer skills in vehicle control. beginner drivers are often subject to influence of opinions of other car occupants, they often drive vehicles in poor technical condition and tend to drive high-speed powerful vehicles and thus are exposed to a false feeling of safety. statistical research confirms that young drivers are referred as the most dangerous group. they, more often than others, exceed the speed limit, overtake in sections with low visibility road sections or when not allowed and tend to hit the red traffic light signal. in this group there are also individuals who tend to drive more often under the influence of alcohol and other drugs. young male drivers tend to take more risky maneuvers, rather than young women who tend to avoid dangerous situations, by extent and frequency. young males are also less attentive than females of the same age category. it is, of course, worth noticing that percentage of female drivers in general is still lower than that of male, women drive less in general (kilometers-wise)[7, 8]. driving school instructors in many european countries believe that nowadays, young drivers behave more hazardously on the road, not realizing the potential threat of accidents and their consequences. speeding is associated with masculinity. nowadays, young drivers’ attention, compared to the one in the past, is affected by many disturbing factors such as advertising and bill-boards, loud car acoustic systems, mobile phones etc. driving instructors believe that it may be necessary to reconsider approach to young driver training in terms of study program duration and extension of training programs on driving simulators[9]. 3. traffic accidents with participation of young drivers – case study information for the following study is based on authors’ personal practice of forensic expertise. the expert evaluation of vehicles has been performed, the technical information is given in table 2. the two accidents have some typical features in common. both drivers were from the dangerous group – young drivers. in each case, the accident took place at early morning hours, which makes experts relate to so called "after party" accidents. both were older cars, which may have affected the severity of consequences. the number of passengers at the time of the accident was higher than allowed (6 car occupants in both cases). 4. conclusion dangerous behavior of younger drivers is a reason of many fatal accidents on the roads. one of the important reasons is lack of training for new drivers, which should include both psychological preparation of novice drivers and extensive practical training. also lack of experience is one of the reasons younger drivers overestimate their ability to handle vehicle at high speeds or while turning, neglect the car technical state (old cars) or its power (new cars). this aspect shall be addressed along with other valid methods of accidents rate reduction[10]. references [1] directorate of traffic police 2016. statistics of traffic accidents. online: policie.cz/clanek/statistika-nehodovosti. [2] m. šucha. dopravní psychologie pro praxi: výběr, výcvik a rehabilitace řidičů. praha: grada publishing, 2013. isbn 978-80-247-4113-0. [3] m. šucha. agresivita na cestách. 1. vyd. olomouc: univerzita palackého v olomouci, 2009. isbn 9788024423753. [4] p. bouchner, m. novák, s. novotný, et al. human behavior while driving the road vehicle. prague: faculty of transportation sciences, ctu in prague, department of vehicle technology, 2015. 1st electronic. isbn 978-80-01-05872-5. [5] m. novák, p. bouchner, j. faber, et al. senioři za volantem. praha: department of informatics academy of sciences in czech republic, v.v.i., 2008. isbn 978-80-87136-02-7. [6] s. novotný, p. bouchner, m. novák. agresivní a nebezpečné chování řidičů motorových vozidel v čr – analýza dat z plovoucích vozidel. in: dopravní nehodovost a rizikové chování řidičů motorových vozidel. praha: vyšší policejní škola mv v praze, 2013, pp. 49-59. isbn 978-80-260-5466-5. [7] z. šejdová. riskantní jednání mladých řidičů v české republice. master thesis. palacký university olomouc, 2015. [8] mladí řidiči a dopravní nehodovost. 2007. source: czrso.cz. available on 8.5.2017. [9] goodyear europe, middle east and africa. a report by goodyear europe, middle east and africa (emea). november 2013. 114 policie.cz/clanek/statistika-nehodovosti czrso.cz vol. 12/2017 dangerous driver’s behavior [10] p. bouchner, j. faber, p. hrubeš, et al. road accident reduction. roma: aracne editrice s.r.i., 2010. isbn 978-88-548-3550-4. 115 acta polytechnica ctu proceedings 12:112–115, 2017 1 introduction 2 dangerous driver's behavior 3 traffic accidents with participation of young drivers – case study 4 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0033 acta polytechnica ctu proceedings 7:33–37, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app 1d hermite elements for c1-continuous solutions in second gradient elasticity christian liebold∗, wolfgang h. müller berlin university of technology, chair of continuum mechanics and materials theory, einsteinufer 5, 10587 berlin, germany ∗ corresponding author: christian.liebold@tu-berlin.de abstract. we present a modified strain gradient theory of elasticity for linear isotropic materials in order to account for the so-called size effect. additional material length scale parameters are introduced and the problem of static beam bending is analyzed. a numerical solution is derived by means of a finite element approach. a global c1-continuous displacement field is applied in finite element solutions because the higher-order strain energy density additionally depends on second gradients of displacements. so-called hermite finite elements are used that allow for merging gradients between elements. the element stiffness matrix as well as the global stiffness matrix of the problem is developed. convergence, c1-continuity and the size effect in the numerical solution is shown. experiments on bending stiffnesses of different sized micro beams made of the polymer su-8 are performed by using an atomic force microscope and the results are compared to the numerical solution. keywords: second gradient continuum, size effect, hermite finite elements. 1. introduction materials with intrinsic microor nano-structure can show size-dependent material behavior, which is reflected, e.g., by a stiffer elastic response to external forces when the size of the material body is reduced. a quantitative understanding of a possible size effect in microand submicrostructures is of great importance when modeling microand nano-electro-mechanical systems (mems/nems). driven by miniaturization, the need for saving materials and improving the performance, the requirement of reliability in simulation technologies increases. a size effect is reflected in a different elastic response to external loads if the size of the material body is reduced. experimental validation therefore is given in, e.g., [1–6]. second gradient (sg) continuum theories are used if, in the absence of strain gradients (in uniaxial tensile tests), the elastic behavior of the material is independent of the thickness of the sample [3]. they are applicable to so-called non-simple materials of the gradient type, for example, polymers at a small scale. since conventional continuum theories based on the cauchy continuum are not able to predict size effects, the present work deals with the modified strain gradient theory (msg) developed in, e.g., [7–9]. by employing the macroscopic rotation vector and second order derivatives of the displacement vector, msg-theory is able to account for quantities like rotation and curvature. a generalization of strain gradient continua has been developed by eringen, who proposed “nonsimple materials of gradient type [10]” in order to derive the corresponding higher-order material dependencies in a rational manner. the application of conventional finite element (fe) strategies may lead to inaccurate results, if finite element formulations are used which only fullfill global c0–continuity. the scope of this work is to use fe formulations based on hermite polynomials in order to fulfill global c1–continuity for solutions in the modified strain gradient theory. section 2 presents the strain energy function of the modified strain gradient theory and a one dimensional differential equation is derived by making use of the euler-bernoulli assumptions for slender beams. the finite element formulation based on hermite polynomials is presented in sect. 3 as well as the structure of the element stiffnessand system matrix. a numerical example for a loaded cantilever beam is then given in sect. 4. 2. modified strain gradient theory 2.1. strain energy function the present study starts with one of the three reduced forms of the strain energy densities for small deformations, usg, as postulated by mindlin (1962) [8]. the einstein summation convention on repeated indices is used and spatial partial derivatives in the cartesian coordinate system are denoted by commaseparated indices. mindlin’s second form of linear isotropic strain energy density reads: usg = α1εij εij + α2εkk εmm + β1ηijk ηijk + β2ηiik ηjjk + + β3ηiik ηkjj + β4ηijj ηikk + β5ηijk ηkji , (1) where εij = 1/2(ui,j + uj,i ) denotes the small strain tensor and ηijk = 1/2(uk,ij + uj,ki ) = εkj,i the gradient 33 http://dx.doi.org/10.14311/app.2017.7.0033 http://ojs.cvut.cz/ojs/index.php/app christian liebold, wolfgang h. müller acta polytechnica ctu proceedings of strain. α1 and α2 lead to lamé’s constants and β1,...,5 are five additional higher-order material constants [11]. σij and µijk denote the cauchy stress tensor and the higher-order stress tensor, respectively: σij = ∂usg ∂εij , µijk = ∂usg ∂ηijk . (2) the inclusion of the macroscopic rotation vector ϕi = 1/2�ijk uk,j (�ijk being the levi-civita symbol) leads to a reduction of the additional material constants from five to three. based on fleck & hutchinson (1997) [12] we introduce independent expressions of ηijk and decompose the second order displacement gradient into a symmetric part ηsijk and a remaining part ηrijk (which is not necessarily anti-symmetric): hj hj hj z~ hj �� �� �� �� ηijk ηsijk η r ijk η̄ij χsijχ a ij η (0) ijk η (1) ijkχ a ijεmm,i figure 1. scheme of decomposition ηijk = ηsijk + η r ijk , η s ijk = 1 3 (uk,ij + ui,jk + uj,ki ) , ηrijk = 2 3 (�ikl η̄lj + �jkl η̄li ) + �kjl η̄li , (3) where η̄ij = ϕj,i is the gradient of rotation, which we decompose into its symmetric and anti-symmetric parts, χsij and χaij , respectively: χaij = 1 2 (ϕi,j −ϕj,i ) , χ s ij = 1 2 (ϕi,j + ϕj,i ) . (4) the tensor ηsijk is further decomposed into its spherical and deviatoric part, η(0)ijk and η (1) ijk , cf., fig. 1 . the quantity η(0)ijk is related to χ a ij and the dilatation gradient εmm,i in the following manner: η (0) ijk = 1 5 ( δij η s mmk + δjk η s mmi + δki η s mmj ) ηsmmi = εmm,i + 2 3�iln χ a ln , η (1) ijk = η s ijk −η (0) ijk . (5) as shown in [13, 14] χaij does not influence the strain energy if symmetry of the couple stress tensor µij is assumed. as a consequence the linear isotropic strain energy density for non-simple materials of gradient type reads: umsg = 2gεij εij + λεkk εii + 2g`20εmm,i εkk,i + + 2g`21η (1) ijk η (1) ijk + 2g` 2 2χ s ij χ s ij , (6) where λ and g are lamé’s constants, whereas `0,`1 and `2 denote additional material length scale parameters. these parameters are chosen to be of the dimension of a length and squared in order to guarantee a positive definite problem in the energy minimization. the normalization of the higher-order terms by g is arbitrary. 2.2. governing euler-bernoulli differential equation by using the displacement field of an eulerbernoulli beam of length l, thickness t, crosssection a, young’s modulus e, and second moment of inertia i cf., fig. 2, ux = −z dw(x) dx , uy = 0 , uz = w(x), (7) l, ei, a w(x) dw(x) dx f xy z figure 2. bending line of a cantilever where w(x) characterizes the bending line and f denotes the force at the free end, eqs. (3)–(6) lead to the strain energy density of the body. by applying the principle of virtual work to the variation of the function w(x), a triple integration by parts leads to the ordinary differential equation (ode) of the problem [15]: s d4w dx4 −k d6w dx6 = 0 , ∀ x ∈ [0,l] . (8) the resulting ode is of rank six, where s = ei(1 + 4308225 `2 t2 ) and k = 75ei` 2 (9) are given constants, if a rectangular cross-section and `0 = `1 = `2 = ` is assumed. 3. finite element implementation 3.1. weak form and discretization a multiplication of eq. (8) with a test function v(x) and triple integration by parts yields in the following weak integral form of the ode: l∫ 0 ( s d2w dx2 d2v dx2 + k d3w dx3 d3v dx3 ) dx = 0 . (10) now the galerkin method is applied to discretize the interval [0,l] by using linear combinations of independent base functions φα for each element e = 1, . . . ,n, which are uniformly defined on the element’s coordinate 0 ≤ ζ ≤ 1: ve(ζ) = ∞∑ β=1 φβ , w e(ζ) = ∞∑ δ=1 ceδφδ . (11) the general element function of this method is derived by inserting eq. (11) in eq. (10), evaluated on the domain ζ: 1∫ 0 ∞∑ δ=1 ∞∑ β=1 ( sceδ d2φδ dζ2 d2φβ dζ2 + kceδ d3φδ dζ3 d3φβ dζ3 ) dζ = 0 . (12) 34 vol. 7/2017 c1-continuous solutions in second gradient elasticity 3.2. hermite elements for global c1–continuity in order to solve the present weak form of the problem, the base functions need to be differentiable three times at least. in order to guarantee that information of second gradients of the trial function w(x) is included, we ask for global continuity of the first derivative of w(x), which is called c1–continuity of the solution. therefore, we choose a linear combination of four different hermite polynomials as follows: h1 = 2ζ3 − 3ζ2 + 1 , h2 = ζ3 − 2ζ2 + ζ h3 = −2ζ3 + 3ζ2 , h4 = ζ3 − ζ2, (13) h1 h2 h3 h4 0 ≤ ζ ≤ 1 figure 3. four hermite polynoms per element which allow to influence the first derivative at each boundary of the element independently. the unit element now is a linear combination of the four hermite polynomials: ve(ζ) = 4∑ β=1 hβ , w e(ζ) = 4∑ δ=1 ceδhδ (14) where ceδ are four constants per element. 3.3. structure of the element stiffnessand system matrix by using hermite elements as base functions eq. (12) can be written in the following manner: 1∫ 0 4∑ δ=1 4∑ β=1 ( sceδ d2hδ dζ2 d2hβ dζ2 +kceδ d3hδ dζ3 d3hβ dζ3 ) dζ = 0. (15) the element stiffness matrix ke reads: keδβ = 1∫ 0 ( s d2hδ dζ2 d2hβ dζ2 + k d3hδ dζ3 d3hβ dζ3 ) dζ =   ke11 k e 12 k e 13 k e 14 ke22 k e 23 k e 24 sym. ke33 ke34 ke44   = const. (16) and is a symmetric 4×4 matrix with constant components that depend on the material data, beam geometry, and on the length of the element. in this approach we choose equidistantly distributed elements. by special assembling of the system matrix it is now possible to connect the boundary values of an element to its neighbors and, in addition, to bring together the first derivatives in between. this will be achieved by the following structure of the system matrix k: k =   k1δβ + k2δβ + k3δβ + k···δβ + knδβ   . (17) due to the summation of the last entries, the specific components of neighboring elements are treated as equivalent. this refers to both, the components that represent direct node values as well as to the corresponding derivatives. the solution of the entire beam bending problem is now approximated by a linear system of equations: kc = f ⇒ k\f = c , (18) where c is the vector of element coefficients and f is a right hand side, in which the external force f is implemented (here, situated at the very last node n). the system is solved using the ”backslash” operator of the commercial program matlab (lu decomposition, etc.). 4. numerical example as an example we study bending of a cantilever beam, fixed at the left hand side and loaded with a single force at the free end, as presented in fig. 2 . material and geometry data is given in tab. 1 . the material length scale parameter ` is chosen according to a literature value [3] and to measurements performed by the authors with an atomic force microscope [16, 17]. t w l e `0 = `1 = `2 f 100 µm 2t 20t 1.44 gpa 17.6 µm 1 µn table 1. dimensions, loads and material parameters the boundary condition for the fixation has been implemented by canceling the first and second row and column of the system matrix eq. (17), since, regarding the first node of the first element, the current values of deflection and the derivative of the deflection line are equal to zero. 35 christian liebold, wolfgang h. müller acta polytechnica ctu proceedings er r ×10−5 0 5 10 15 20 0 40 80 120 160 number of elements n figure 4. convergence the test for convergence with respect to higher numbers of elements n yield positive results (cf. fig. 4). the relative error of tip deflections err= wn/w1 −1 per increase of elements achieves values below 0.001 h. lkmberlin university of technologyfaculty v – mechanical engineering and transport systems institute of mechanics chair of continuum mechanics and materials theory prof. dr. rer. nat. w. h. müller table: dimensions, loads and material parameters 𝑡 𝑊 𝐿 𝐸 ℓ0 = ℓ1 = ℓ2 𝐹 100 µm 2𝑡 20𝑡 1.44 gpa 17.6 µm 1 µn 𝑤 [m m ] 𝑥/𝐿 d𝑤 d𝑥 𝑥/𝐿 d2 𝑤 d𝑥2 𝑥/𝐿 d3 𝑤 d𝑥3 𝑥/𝐿 dr.-ing. c. liebold | exnum 2016, liblice | 1d example 9/10 figure 5. deflection line the deflection line of the beam solution is presented in fig. 5 . five elements are shown so that a closer look at the solution of each element becomes possible, which consists of reconstructed trial functions, cf. eq. (14)2. figure 6 presents a plot of the first derivative of the reconstructed trial functions per element and clearly demonstrates continuity in between neighboring elements and therefore c1–continuity. lkmberlin university of technologyfaculty v – mechanical engineering and transport systems institute of mechanics chair of continuum mechanics and materials theory prof. dr. rer. nat. w. h. müller table: dimensions, loads and material parameters 𝑡 𝑊 𝐿 𝐸 ℓ0 = ℓ1 = ℓ2 𝐹 100 µm 2𝑡 20𝑡 1.44 gpa 17.6 µm 1 µn 𝑤 [m m ] 𝑥/𝐿 d𝑤 d𝑥 𝑥/𝐿 d2 𝑤 d𝑥2 𝑥/𝐿 d3 𝑤 d𝑥3 𝑥/𝐿 dr.-ing. c. liebold | exnum 2016, liblice | 1d example 9/10figure 6. first derivative of the solution as expected the global solution is not c2–continuous. this is confirmed by the jumps between the elements in their second derivative in fig. 7. such jumps are also present in the third derivative in fig. 8, but invisibly small. lkmberlin university of technologyfaculty v – mechanical engineering and transport systems institute of mechanics chair of continuum mechanics and materials theory prof. dr. rer. nat. w. h. müller table: dimensions, loads and material parameters 𝑡 𝑊 𝐿 𝐸 ℓ0 = ℓ1 = ℓ2 𝐹 100 µm 2𝑡 20𝑡 1.44 gpa 17.6 µm 1 µn 𝑤 [m m ] 𝑥/𝐿 d𝑤 d𝑥 𝑥/𝐿 d2 𝑤 d𝑥2 𝑥/𝐿 d3 𝑤 d𝑥3 𝑥/𝐿 dr.-ing. c. liebold | exnum 2016, liblice | 1d example 9/10 figure 7. second derivative of the solution lkmberlin university of technologyfaculty v – mechanical engineering and transport systems institute of mechanics chair of continuum mechanics and materials theory prof. dr. rer. nat. w. h. müller table: dimensions, loads and material parameters 𝑡 𝑊 𝐿 𝐸 ℓ0 = ℓ1 = ℓ2 𝐹 100 µm 2𝑡 20𝑡 1.44 gpa 17.6 µm 1 µn 𝑤 [m m ] 𝑥/𝐿 d𝑤 d𝑥 𝑥/𝐿 d2 𝑤 d𝑥2 𝑥/𝐿 d3 𝑤 d𝑥3 𝑥/𝐿 dr.-ing. c. liebold | exnum 2016, liblice | 1d example 9/10figure 8. third derivative of the solution e ∗ [g p a] thickness t [mm] figure 9. size effect for decreasing thicknesses the sensitivity of the implemented numerical model of the modified second gradient continuum was tested for incorporation of size effect by decreasing the thickness values, while all other dimensions were scaled in parallel according to tab. 1 . the solution of the tip deflection was used in a calculation of an equivalent young’s modulus, e∗, with a euler-bernoulli analytical beam model of the conventional cauchy continuum. as plotted in fig. 9, a size effect is clearly present. for small thicknesses the equivalent young’s modulus is calculated to be about three times larger than the usual one. 36 vol. 7/2017 c1-continuous solutions in second gradient elasticity 5. conclusions based on a modified strain gradient theory for linear isotropic materials we have evaluated the eulerbernoulli beam assumptions to derive the ordinary differential equation of the problem. the weak form was derived using test functions. third order hermite polynomials have been summed up for an element formulation: these are able to access and influence first derivatives between the elements. the element stiffness and system matrix have been formulated for the problem of a cantilever beam. the resulting system of linear equations was solved numerically to determine the coefficients of the trial functions for each element. the global solution was reconstructed element-wise and first, second and third order derivatives have been calculated. it was demonstrated, that this procedure guarantees c1–continuity of the solution. this is considered to be an important quality in second gradient theories. furthermore, the size effect, measured by the authors in previous work, could be reproduced by using similar material as well as geometry data. this provides the impetus for the development of a three dimensional finite element formulation of a strain gradient continuum on the basis of hermite elements. references [1] w. j. poole, m. f. ashby, n. a. fleck. micro-hardness of annealed and work-hardened copper polycrystals. scripta materialia 34(4):559–564, 1996. [2] x. h. guo, d. n. fang, x. d. li. measurement of deformation of pure ni foils by speckle pattern interferometry. mechanical engineering 27(2):21–25, 2005. [3] d. c. c. lam, f. yang, c. m. chong, et al. experiments and theory in strain gradient elasticity. j mech phys sol 51(8):1477–1508, 2003. [4] a. w. mcfarland, j. s. colton. role of material microstructure in plate stiffness with relevance to microcantilever sensors. journal of micromechanics and microengineering 15(5):1060–1067, 2005. [5] s. cuenot, c. fretigny, s. demoustier-champagne, b. nysten. surface tension effect on the mechanical properties of nanomaterials measured by atomic force microscopy. physical review b 69:01–05, 2004. [6] x.-f. li, b.-l. wang, k. y. lee. size effect in the mechanical response of nanobeams. j of adv research in mech eng 1(1):4–16, 2010. [7] r. toupin. elastic materials with couple-stresses. arma 11:385–414, 1962. [8] r. mindlin. effects of couple-stresses in linear elasticity. arma 11:415–448, 1962. [9] r. d. mindlin, n. n. eshel. on first strain-gradient theories in linear elasticity. international journal of solids and structures 4:109–124, 1968. [10] a. eringen. nonlocal continuum field theories. springer-verlag, new york 2010. [11] m. lazar. irreducible decomposition of strain gradient tensor in isotropic strain gradient elasticity. zamm z angew math mech 2016. doi:10.1002/zamm.201500278. [12] n. a. fleck, j. w. hutchinson. strain gradient plasticity. advances in applied mechanics, academic press, new york 33:295–361, 1997. [13] c. liebold, w. h. müller. measuring material coefficients of higher gradient elasticity by using afm techniques and raman-spectroscopy. advanced structured materials, springer berlin heidelberg 22:255–271, 2013. [14] f. yang, c. m. chong, d. c. c. lam, p. tong. couple stress based strain gradient theory for elasticity. international journal of solids and structures 39(10):2731–2743, 2010. [15] s. kong, s. zhou, z. nie, k. wang. static and dynamic analysis of micro beams based on strain gradient elasticity theory. international journal of engineering science 47:487–498, 2009. [16] c. liebold, w. h. müller. applications of strain gradient theories to the size effect in submicrostructures incl. experimental analysis of elastic material parameters. bulletin of ticmi 19(1):45–55, 2015. [17] c. liebold, w. h. müller. comparison of gradient elasticity models for the bending of micromaterials. elsevier, computational materials science 116:52–61, 2016. 37 http://dx.doi.org/10.1002/zamm.201500278 acta polytechnica ctu proceedings 7:33–37, 2017 1 introduction 2 modified strain gradient theory 2.1 strain energy function 2.2 governing euler-bernoulli differential equation 3 finite element implementation 3.1 weak form and discretization 3.2 hermite elements for global c1–continuity 3.3 structure of the element stiffnessand system matrix 4 numerical example 5 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0086 acta polytechnica ctu proceedings 11:86–90, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app analysis of competitiveness of railway passanger transport between rakovník and prague rudolf vávra∗, vít janoš department of logistics and management of transport, faculty of transportation sciences, ctu in prague, czech republic ∗ corresponding author: vavrarud@fd.cvut.cz abstract. this paper deals with the issue of railway connection between rakovník and prague. initially, there is made an analysis of the transport supply and analysis of modal split. additionally, an operation concept is suggested in order to increase the attractiveness of the railway passenger transport, afterward, the transport possibilities, which can be accomplished by this approach, are evaluated. keywords: modal split, transport supply, logit model, operation concept, periodic timetable, travel time, public transport. 1. introduction the article deals with examining the competitiveness of the railway connection between prague and rakovník. rakovník (16081 inhabitants) is a town located in the western part of středočeský kraj (region of middle bohemia), outside the suburban area of prague, but some properties are similar to a suburban area – has strong daily transport relations due prague (according to the census from 2011 about 738 people daily). rakovník to prague does have good bus connection, but due to the existence of the railway connection in this session, and the advantages of rail transport through the continuous passage of agglomeration areas should address the issue of the attractiveness of this mode of transport. among solved railway lines are included (line numbers by railway timetable sždc): • 120 praha – kladno – rakovník, • 171 praha – beroun, • 173 praha – rudná u prahy – beroun, • 174 beroun – rakovník. the aim is to examine the competitiveness of rail passenger transport in relation rakovník – prague, using the current infrastructure, which includes: • an analysis of the modal split for the current state, • determination of the theoretical travel times possibilities of the train and comparing them with the travel times according to the current state, • proposal schematic operational concept in term of technological possibilities, • comparing travel times of created operational concept with travel times according to the current state and theoretically achievable travel times, • determination of the modal split for the proposed operational concept and measurement of the difference in the modal split of proposed operational concept from the current situation. 2. methods to determine the share of individual vehicles on the modal split is used logit model, namely a two-step nested logit model: pi = e µvi∑ j∈dp e µvj where pi means the probability of choice of transport mode i, e is euler’s number, dp is the set of all considered transport modes, µ means the elasticity of choice of transport mode on the basis of changes in generalized costs and vi means generalized costs expended for the journey by transport mean i. [1] nested logit model is phasing out the calculation in two phases. the first phase determines modal split between individual transport and public transport. the second phase examines then modal split of individual transport modes within public traffic. 2.1. first phase in the first phase it is necessary to determine the generalized cost of individual car transport and public transport. within public transport there is selected that transport mode of public transport, that is more favourable for the journey. generalized costs of individual car transport (viad) are calculated using the formula 1, where dkm is length of the journey (in km), niad are costs for 1 km of the journey, oiad is the average occupancy of the vehicle, diad čas is travel time (in hours), wčas is the perceived value of time (in czk/hour) and diad dost is an indicator of time availability in hours (includes attendance from origin of the journey to the car, time attendance from the car to final destination, time for finding a parking space, etc.). generalized costs of public transport (vv d) are calculated using the formula 2, where dv d čas is travel 86 http://dx.doi.org/10.14311/app.2017.11.0086 http://ojs.cvut.cz/ojs/index.php/app vol. 11/2017 analysis of competitiveness of railway passanger transport viad = dkm ∗ niad oiad + diad čas ∗ wčas + diad dost ∗ wčas (1) vv d = ( dv d čas ∗ wčas + dv d dost ∗ wčas + jv d + iv d,1/2 ∗ wčas ) ∗ cv d−iad (2) time (in hours), wčas is the perceived value of time (in czk/hour), dv d dost is an indicator of time availability in hours (time attendance from origin of the journey to the stop/station of public transport, time attendance from the stop/station to final destination), jv d is fare (in czk), iv d,1/2 is half the interval between links (in hours) and cv d−iad is the calibration constant of mode choice. [1] the result of the first phase and are therefore values piad and pv d indicating the probability of the use of individual car traffic and public transport in the appropriate relation. 2.2. second phase in second phase there is necessary to determine generalized costs (vi) for relevant modes i of public transport. determination of these costs is performed according to the formula 3, where di čas is travel time (in hours) using transport mode i, wčas is the perceived value of time (in czk/hour), di dost is time availability of the transport mode i (in hours), ji is fare using transport mode i (in czk) and ii,1/2 is half the interval between links of the transport mode i (in hours). [1] results of second phase are the conditional probabilities pi∨v d of the use of transport mean i under the condition of using public transport. the probability pi of use of the transport mean is thus given by pi = pv d ∗ pi∨v d. 3. results and discussion 3.1. analysis of supply connection during the timetable 2015/2016 bus connection from rakovník to prague was during the timetable 2015/2016 accounted by lines 310610, 310612, 310613, which are operating between rakovník bus station and stop prague hradčanská and further by lines 310270, 310700, 310710, which are operating between rakovník bus station and stop prague zličín. average aggregate interval between individual links reaches during peak hours 10 minutes. shortest travel times between rakovník and hradčanská are about 55 to 75 min. the shortest travel times between rakovník and zličín are about 40 to 55 min. by train there is possible to go either using line 120 through kladno or using the lines 174 and 171 (eventually 174 and 173) through beroun. on the line 120 are operated in 120 minutes period fast trains praha – rakovník. connection rakovník to prague is complemented by a combination of fast regional trains praha – kladno (in 120 minutes period, between praha – kladno in leading with fast trains) and passenger trains kladno – rakovník (operated as the 120 minute period) with bond in the station kladno. the connection is also completed in peak hours of working days with one pair of fast regional trains rakovník – prague. these connections are operated in periodic timetable. on the line 174 are operated passenger trains beroun – rakovník, but without a fixed interval between links (i.e. not according periodic timetable). their average interval during peak hours of working days is around 60 minutes. in beroun is then formed a bond in direction prague. 3.2. analysis of modal split in timetable 2015/2016 for the analysis of the modal split between prague and rakovník were selected 3 relations. in rakovník is always considered the town centre. in prague, with respect to potential target for daily commuting, were selected these points: vítězné náměstí, anděl, václavské náměstí. to analyze the modal split was used nested logit model described in chapter 2. to determine values dkm and diad čas was used free route planner on website mapy.cz. the value niad is determined as the sum of the fuel cost per 1 km away (at an average consumption of 7 litres per 100 km) and other costs per 1 km of ride (related to the operation and acquisition of the vehicle). rates are taken from the decree no. 328/2014 coll., so niad = 6, 213 czk/km. for the average occupancy of the vehicle according to chmelík [1] was calculated with the value oiad = 2. for variable value diad dost is selected diad dost = 0, 25 h (5 minutes attendance from the journey origin to the vehicle, 5 minutes attendance from the car to the final destination and 5 minutes time on finding a parking space and parking). in the case of public transport is taken into account connection by bus and by train in the working day (on september wednesday as an average traffic day) in the direction from prague to rakovník, with the departure of the train / bus from prague in the time between 14:00 and 18:00. values di čas and ii,1/2 are based on supply of connections, with ii,1/2 = 2ps, where ps means the number of considered connections in the solved time period. the value of the availability di dost of transport mean i is determined by the route planner and includes: 87 rudolf vávra, vít janoš acta polytechnica ctu proceedings vi = di čas ∗ wčas + di dost ∗ wčas + ji + ii,1/2 ∗ wčas (3) figure 1. current modal split. • attendance of the reference point in prague to the appropriate stop/station of transport mean i , or to the stop of prague urban transport, • attendance from the bus station in rakovník, or from the railway station rakovník to the town centre, • additional reserve of 5 min. fare ji is determined from the carriers’ tariffs. for bus services is being considered at ordinary fare payment smart card, for rail services with custom fare in 25, at the prague urban transport then using annual coupon for an adult. perceived time value is calculated according to chmelík [1] as 0.24 multiple of the average gross hourly wages, while gross monthly wage czk 26480 and 168 working days in 2016 and 8 working hours per working day brings wčas = 37, 82857 czk/hour. values µ and cv d−iad were chosen after chmelík [1], therefore cv d−iad = 2, 2, µ = −0, 0031 for the first phase of the model and µ = −0, 0123 for second phase of the model. the resulting values for modal split for the current state shows figure 1. 3.3. determination of the theoretically achievable travel times theoretically achievable travel time was determined on the basis of completely free infrastructure capacity. there are considered no stops and therefore is neglected a number of restrictive conditions. to determine theoretically achievable rail travel time was used software fbs. theoretically achievable travel times were determined for followed relations: • praha masarykovo n. – rakovník, • praha-dejvice – rakovník, • praha hl.n. – rakovník, • praha-smíchov – rakovník. theoretically achievable travel times in these relations, as well as its comparison with the shortest travel time by train timetable 2015/2016, the shortest travel time by bus (or in combination with prague urban transport) and travel time by car, shows the graph in figure 3. 3.4. the proposal of schematic operational concept by the real technological possibilities despite the temporal lack of competitiveness of rail passenger services between prague – rakovník, which is evident from the graph in figure 3, was for solved lines created schematic operational concept by the real technological possibilities, that would have the effect of increasing the value of modal split compared to current situation. this can be achieved in the following ways: • shorting the travel times, • shorting the interval between links, • increasing number of direct rail links between rakovník and prague, • using principles of integrated periodic timetable (ipt). for creating the operational concept was used fbs. on the line 120 there is proposed hourly operation of fast trains between rakovník – prague, instead of the current two-hour interval. the line reaches the broader ipt node in minute 00 in rakovník. regional trains kladno – rakovník are kept in a two-hour interval with bond to passenger trains prague – kladno in kladno. on lines 171 and 174 is proposed introducing of fast regional trains prague – beroun – rakovník, which would be operated in 120 minutes interval. on the line 174 is also proposed to introduce two lines of trains, both kept in interval 120 minutes. ipt node in minute 00 in rakovník achieves, however, only one of them. on the 173 was taken into account the introduction of peak 20 minutes interval between prague smíchov and nučice station. for connection between rakovník – prague this line does not have much significance. the proposed operational concept (for peak hours) is displayed in the form of interval graphics in the figure 2. the graph in figure 4 shows similar information as the graph in figure 3, to which are added the shortest rail travel times from proposed operational concept. 3.5. determination of the modal split for the proposed operational concept for new proposed operational concept was calculated modal split again. it was used the same method of calculation as in section 3.2. 88 vol. 11/2017 analysis of competitiveness of railway passanger transport figure 2. interval graphic of the proposed operational concept. 89 rudolf vávra, vít janoš acta polytechnica ctu proceedings figure 3. comparison of theoretically achievable rail travel time. figure 4. comparison of theoretically achievable rail travel time, including shortest rail travel times from proposed operational concept. modal split for the proposed operational concept is shown in the graph in figure 5. the graph in figure 6 shows the modal shift of the proposed scheme compared to the level in force during the timetable 2015/2016. 4. conclusions the aim was to examine the competitiveness of rail passenger transport between rakovník and prague. of the share of modal split of the relevant transport modes in relation rakovník – prague, it was found, that the share of train is today in comparison with individual car traffic and the bus very low. then there were determined for the train relation rakovník – prague theoretical possibilities of achievable travel time, and it was found, that even when neglecting a number of restrictive conditions, the train could not be between prague and rakovník time-competitive, even compared to the bus, let alone compared to individual car transport. to calculate the modal-split was used nested-logit model. then was created an operational concept aimed at increasing the attractiveness of passenger rail transport in relation rakovník – prague. for the proposed operational concept has been found, how the train travel time could be closer to the theoretically achievable travel time. for timetable construction was used fbs. it was found, which transportation options may be achieve due to new operational concept and for this solution were determined new values of the modal split. after comparing new values with the previous (original) the results shows, that despite the proposed offer in passenger rail service between prague and figure 5. modal split in relations prague – rakovník using the proposed operational concept. figure 6. modal shift in relations prague – rakovník using the proposed operational concept. rakovník was not achieved appreciable improvement, and value of modal split for the train increased by only 0.42 %. the result of examination is that possibilities of contemporary rail infrastructure are so limited, that changes in timetable concept and offer of more connection does not allow significant improvements. so it is confirmed, that between rakovník and prague is not possible to achieve fundamental changes in the modal split without significant changes on to the rail infrastructure. references [1] j. chmelík. assessments of modal split in long-distance passenger transport [pdf]. review of economic perspectives – národohospodářský obzor, 2015. [cit. 2016-08-01]. 90 acta polytechnica ctu proceedings 11:86–90, 2017 1 introduction 2 methods 2.1 first phase 2.2 second phase 3 results and discussion 3.1 analysis of supply connection during the timetable 2015/2016 3.2 analysis of modal split in timetable 2015/2016 3.3 determination of the theoretically achievable travel times 3.4 the proposal of schematic operational concept by the real technological possibilities 3.5 determination of the modal split for the proposed operational concept 4 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2018.18.0044 acta polytechnica ctu proceedings 18:44–47, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app impact testing of ordnance gelatine under moderate strain rate conditions tomáš doktor∗, petr zlámal, jan šleichrt, tomáš fíla, daniel kytýř czech technical university in prague, faculty of transportation sciences, konviktská 20, praha 1, czech republic ∗ corresponding author: doktor@fd.cvut.cz abstract. an experimental study on energy absorption capabilities and strain rate sensitivity of ordnance gelatine was performed. strain energy density under quasi static compression and moderate strain rate impact tests was compared. in the study two types of material were tested, bulk ordnance gelatine and polymeric open-cell meshwork filled with ordnance gelatine. from the results a significant strain-rate effect was observed in terms of ultimate compressive strength and strain energy density. in comparison of the deformation behaviour under quasi static conditions and drop weight test the difference was very significant, however slight increase in both strength and strain energy density was observed even between different impact energies and velocities during the impact testing. the peak acceleration was significantly reduced in polymer meshwork filled by gelatine in comparison to the bulk gelatine. keywords: ordnance gelatine, strain energy density, impact testing. 1. introduction in the engingeering applications in the field of ballistic and blast protection, passive safety systems in vehicles (automotive, railway vehicles) the impact energy dissipation is among the crucial properties [1]. one of the possible solutions to achieve such capabilities is to utilise cellular solids in conjunction with strain rate sensitive filling [2]. in this study experimental investigation on strain rate sensitivity at moderate strain rates as well as the interaction of the cellular solid and viscous filling is described. as filling material ordnance gelatine was selected due to a flawless preparation procedure and availability. the ordnance (ballistic) gelatine is widely used as tissue simulant for assessment of damage of tissues induced by bullets [3]. however strain rate sensitivity under compressive loading was reported as well [4]. deformation response of ballistic gelatine blocks (plain and reinforced with polymeric meshwork) under moderate impact loading is described in this paper in terms of peak acceleration during the impact as well as total kinetic energy dissipation. 2. materials and methods 2.1. sample preparation for both impact and quasi-static testing two types of material were used, (i) bulk ordnance gelatine and (ii) open-cell polymeric meshwork filled with ordnance gelatine. the gelatine was prepared according to the reports of jussila [5]. as a base, 260 bloom beef gelatine (remi mb, ltd., czech republic) was used. the gelatine powder was poured into warmed water (temperature was 45 °c) in the second group the gelatine was reinforced with polymeric open-cell meshwork ft-s10fr (foam techniques, ltd, united kingdom) with tetrakaidecahedral cells, mean cell size 6 mm, mean strut thickness 0.6 mm. the dissolved gelatine after homogeneous mixing was poured into the meshwork. the samples from both batches (both pure and reinforced) were cured for 24 h in room temperature and subsequently stored for 24 h in refrigerator. after this curing the blocks were cut into samples of desired size. in the quasi static loading conditions the size of samples was 25 × 25 × 50 mm, and samples’ size for impact testing was approximately 60 × 50 × 50 mm. dimensions of the samples were limited by diameter of used loading plates. to ensure a better focus during optical observation captured face of each sample was sprayed granit paint to obtain random pattern. 2.2. quasi static tests the quasi static tests were performed using electromechanic uniaxial loading device instron 6530 (instron, inc., usa). the compressive tests were displacement driven with loading rate 3 mm · min−1, which corresponded to strain rate 0.001 s−1. maximum displacement was set to 30 mm. 2.3. impact tests instrumentation the moderate strain rate compression tests were performed by an in-house drop tower developed at ctu fts [6]. the drop weight is released by electromagnetic member and then guided on steel rods and induces the impact on a sample places at a stable plate. the drop-tower setup was instrumented by triaxial accelerometer (egcs3, te connectivity, ltd., 44 http://dx.doi.org/10.14311/app.2018.18.0044 http://ojs.cvut.cz/ojs/index.php/app vol. 18/2018 impact testing of ordnance gelatine figure 1. drop tower setup (1 motorized dropweight lift, 2 trigger inductive sensor, 3 guide rods, 4 electromagnetic release member, 5 drop-weight guide frame, 6 accelerometer, 7 drop-weight). usa) with loading capacity ±1000g and impact force transducer (200c20, pcb piezoelectronics, usa) with loading capacity 89 kn. both accelerometer and force transducer were connected to read out electronics ni9234 (national instruments, inc., usa) providing sampling rate 51200 samples per second. for the observation of deformation process a high speed camera (idt nx3, usa) was employed. the camera provided frame rate 2895 to 3310 fps at resolution 768 × 1312 px and the exposure time varied from 89 to 115 µs. for illumination of the loading scene pair of flash illuminators veritas constellation 60 were used. the impact test setup is depicted in figure 1 and figure 2. 2.4. impact test procedure the impact tests were carried out in three arrangements with different initial height and mass of the impactor. in each arrangement a pair of samples from the group a (bulk gelatine) as well as from group b (polymer meshwork filled with gelatin) were tested. the initial heights and impactor masses are listed in table 1 with corresponding impact energy, velocity and strain rate. the impact velocity was calculated using the initial heigt and gravity constant based on conservation of sum of kinetic and potential energy. figure 2. drop tower setup (1 specimen, 2 impactor, 3 flash-illumination, 4 high-speed camera). 0 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0.5 tr u e s tr e s s [ m p a ] true strain [-] stress strain diagrams at quasi static compression a, 0.001/s b, 0.001/s figure 3. stress-strain curves of quasi-static tests. 3. results from the quasi static loading test stress-strain curves which are depicted in figure 3 were calculated based on the measured force f and cross-head displacement u by formulae σeng = f a0 and �eng = u l0 , where a0 is specimen cross-section and l0 is the initial gauge length. the engingeering stress σeng and strain �eng were then converted to logarithmic strain � and true stress σ using formulae � = ln�eng and σ = σeng(1 + �eng). to evaluate the drop tests stress-strain diagrams were created from the measured acceleration and force. while the stress is measured directly by force transducer, strain calculation is based on double time integration of the vertical component of acceleration. 45 t. doktor, p. zlámal, j. šleichrt et al. acta polytechnica ctu proceedings arrangement initial height drop weight impact velocity impact energy [m] [kg] [ms-1] [j] 1 1.00 4.495 4.4 44.1 2 1.50 4.495 5.4 66.1 3 1.50 6.504 5.4 95.7 table 1. impact test parameters. 0 2 4 6 8 10 12 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 tr u e s tr e s s [ m p a ] true strain [-] stress strain diagrams group a 73/s, 44.1 j 73/s, 44.1 j 90/s, 66.1 j 90/s, 66.1 j 90/s, 95.7 j 90/s, 95.7 j figure 4. stress-strain curves of drop tests of group a (ordnance gelatine). 0 1 2 3 4 5 6 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 tr u e s tr e s s [ m p a ] true strain [-] stress strain diagrams group b 88/s, 44.1 j 88/s, 44.1 j 108/s, 66.1 j 108/s, 66.1 j 108/s, 95.7 j 108/s, 95.7 j figure 5. stress-strain curves of drop tests of group b (polymer meshwork filled with ordnance gelatine). comparison of the stress-strain curves is depicted in figure 5. for the assessment of dissipated impact energy strain energy density λ was computed using formula λ = ∫ �max 0 σ d� the obtained values are presented in table 2. moreover for visual evaluation of the deformation behaviour the captured loading scene was used (selected series of loading scenes in distinct loading steps is depicted in figure 6). no. / strain energy peak group density [jcm-3] acceleration 1 / a1 1.04503 234.92g 2 / a1 1.00956 209.23g 3 / b1 0.57674 95.01g 4 / b1 0.74435 94.07g 5 / a2 1.12309 210.53g 6 / a2 0.98409 231.10g 7 / b2 0.55441 149.72g 8 / b2 0.79479 147.84g 9 / a3 1.42620 199.30g 10 / a3 1.69909 223.24g 11 / b3 1.22701 152.42g 12 / b3 1.24369 148.73g table 2. impact test results. 4. conclusions an experimental study on deformation behaviour of ordnance gelatine and polymeric open-cell meshwork. from the results a significant strain-rate effect was observed in terms of ultimate compressive strength and strain energy density. in comparison of the deformation behaviour under quasi static conditions and drop weight test the difference was very significant, however slight increase in both strength and strain energy density was observed even between different impact energies and velocities during the impact testing. the comparison between tested types of materials shows a significant reduction of the peak acceleration during the impact while the strain energy density decreased only slightly. this behaviour may be explained by the flow of the filling material thru the meshwork during the impact. the obtained results will provide a base for further modelling of the impact energy dissipation as well as for optimisation of the filling material used in cellular solids. acknowledgements the research was supported by the czech science foundation (research project no. 15-15480s) and by 46 vol. 18/2018 impact testing of ordnance gelatine figure 6. loading scene of selected sample (group a3) in several loading steps (time step between captured images is 0.89ms). czech technical university in prague (research projects sgs17/148/ohk2/2t/16 and sgs18/155/ohk2/2t/16). all the support is gratefully acknowledged. references [1] p. qiao, m. yang, f. bobaru. impact mechanics and high-energy absorbing materials: review. journal of aerospace engineering 21(4):235–248, 2008. doi:10.1061/(asce)0893-1321(2008)21:4(235). [2] t. doktor, p. zlamal, t. fila, et al. properties of polymer-filled aluminium foams under moderate strain-rate loading conditions. materiali in tehnologije 49(4):597–600, 2015. doi:10.17222/mit.2014.195. [3] j. breeze, m. midwinter, d. pope, et al. developmental framework to validate future designs of ballistic neck protection. british journal of oral and maxillofacial surgery 51(1):47–51, 2013. doi:10.1016/j.bjoms.2012.03.001. [4] j. kwon, g. subhash. compressive strain rate sensitivity of ballistic gelatin. journal of biomechanics 43(3):420–425, 2010. doi:10.1016/j.jbiomech.2009.10.008. [5] j. jussila. preparing ballistic gelatine review and proposal for a standard method. forensic science international 141(2-3):91–98, 2004. doi:10.1016/j.forsciint.2003.11.036. [6] o. jirousek, p. zlamal, p. koudelka, t. fila. development and validation of material models for closed-cell metal foam for impact simulation. civil-comp proceedings 106, 2014. 47 http://dx.doi.org/10.1061/(asce)0893-1321(2008)21:4(235) http://dx.doi.org/10.17222/mit.2014.195 http://dx.doi.org/10.1016/j.bjoms.2012.03.001 http://dx.doi.org/10.1016/j.jbiomech.2009.10.008 http://dx.doi.org/10.1016/j.forsciint.2003.11.036 acta polytechnica ctu proceedings 18:44–47, 2018 1 introduction 2 materials and methods 2.1 sample preparation 2.2 quasi static tests 2.3 impact tests instrumentation 2.4 impact test procedure 3 results 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.19.0007 acta polytechnica ctu proceedings 19:7–13, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app scale and serpent two-group cross-section data generation ondřej nováka, ∗, jan frýborta, lubomír sklenkaa, jan rataja, ondřej chválab a czech technical university in prague, faculty of nuclear sciences and physical engineering, department of nuclear reactors, prague, czech republic b university of tennessee at knoxville, department of nuclear engineering, tennessee, usa ∗ corresponding author: ondrej.novak2@fjfi.cvut.cz abstract. in this work, vver fuel assembly is calculated using two different codes serpent and scale. multiplication factor and several two-group macroscopic parameters including scattering cross-section, fission cross-section, total cross-section and diffusion coefficient were studied. in addition to this, fuel isotopic composition dependency on burnup was calculated as a part of this study. obtained data are compared and differences are discussed. both codes provides similar results for all studied parameters. keywords: vver, serpent, vvantage-6, fuel cycle, burnup, scale, temelin nuclear power plant, depletion calculation, homogenized cross-sections. 1. introduction the crucial part of nuclear safety is based on results obtained by calculation. to ensure accurate results, codes are compared with experimental data (benchmarking) and also among each other [8], [6]. in this work, vver fuel assembly depletion is calculated using two different codes monte-carlo code serpent and deterministic sequence triton from scale package. multiplication factor and several two-group macroscopic parameters including scattering cross-section, fission cross-section, total crosssection and diffusion coefficient are calculated and compared. in addition to that, fuel isotopic composition dependency on burnup were calculated. all values are calculated during the fuel depletion in order to cover the whole fuel cycle, which means burnup up to 60.5 mwd/kgu. in addition, several serpent memory optimization settings and their impact on macroscopic data generation were tested. the triton calculation sequence from the scale package [5] represents the traditional approach to fuel depletion calculations and few-group data generation. it is based on deterministic code newt producing data for the well–established origen depletion code. serpent code [9] was selected due to the fact that the code is designed directly for reactor core calculations combining monte-carlo based transport calculations and advanced cram fuel depletion method [10]. serpent and scale codes allow homogenized fewgroup cross-sections and other nuclear constants describing fuel assembly behavior to be produced. these results can be later used in full-core calculations. both codes use different approach to obtain homogenized cross-section and other nuclear data. therefore, comparison provides valuable information about the methods used. scale code has a longer history in this field of application compared to serpent code, thus scale code was selected as the reference code for this study. firstly, the model and the calculation itself is described, then results are presented and discussed. 2. fuel assembly description for this particular study, fuel assembly model was created based on vvantage-6 fuel design [2] (see figure 1). vvantage-6 fuel was developed by westinghouse company for vver-1000 reactors. therefore, it features hexagonal fuel lattice. this fuel type was used in the temelin nuclear power plant reactor for the first 10 years of operation. the fuel assembly without ifba absorber was selected for this study. assembly parameters are listed in the table 1. all fuel pins are identical, with flat 3% enrichment. temperatures used in this study are listed in the table 2. 3. calculation setup two dimensional models of the fuel assembly in serpent code (version 2.1.27) and scale code (version 6.2.1) were created based on the vvantage-6 design specified in table 1. specific power 35.54 w/ghm was used in both calculations. b1-leakage corrected critical spectrum was applied in both calculations as well. in serpent calculation, b1 correction used 238 group energy structure (scale238 [5]). fast group (group 1) was above 0.625 ev and thermal group (group 2) below that for both codes. serpent calculation were conducted with 10000 neutrons per cycle, 400 active cycles and 50 inactive cycles. input cross-section data for serpent calculations were obtained from endf/b-vii.1 library [7]. 7 http://dx.doi.org/10.14311/app.2018.19.0007 http://ojs.cvut.cz/ojs/index.php/app o. novák, j. frýbort, l. sklenka et al. acta polytechnica ctu proceedings figure 1. vvantage-6 fuel assembly [11]. parameter vvantage 6 pin cladding outer diameter [mm] 9.144 pin cladding inner diameter [mm] 8.001 pin cladding thickness [mm] 0.5715 pin cladding material zr-4 (303) fuel pellet outer diameter [mm] 7.844 number of fuel pins in fa 312 pin pitch [mm] 12.75 number of instrumentation tubes 1 tube material zr-4 (303) inner tube diameter [mm] 12.6 outer tube diameter[mm] 11.0 tube thickness [mm] 0.8 number of guide tubes 18 tube material zr-4 (303) outer tube diameter over the absorber area [mm] 12.6 inner tube diameter over the absorber area [mm] 11.0 tube thickness over the absorber area [mm] 0.8 outer tube diameter in the absorber area [mm] 10.41 inner tube diameter in the absorber area [mm] 8.6 tube thickness in the absorber area [mm] 0.905 total assembly weight [kg] 766 uo2 density [g/cm3] 10.412 zr-4 density [g/cm3] 6.53 table 1. fuel assembly parameter summary [3]. 8 vol. 19/2018 scale and serpent two-group cross-section data generation fuel temperature 1005 k soluble boron concentration 525 ppm (3 g/kg) coolant density 0.7169 g/cm3 coolant temperature 578 k table 2. calculation parameters. scale code used the default detailed endf/b-vii.1 252 group library. reflective boundary conditions resulted in infinite lattice calculation for both codes. serpent allows automatic division of fuel pins into radial regions for more realistic depletion calculations. all fuel pins were divided into 5 radial regions and each region was burned separately in this study. in addition, there were unique materials for all fuel pins in the fuel assembly, thus their material compositions were treated separately. scale calculation used 60◦ symmetry to limit the number of depleted materials. fuel pins in symmetrical positions were considered as identical material. no radial regions in fuel pins were defined as there is no burnable absorber. scale and serpent calculation used 51 burn steps with final burnup at 60.5 mwd/kgu1. seven steps were between 0 mwd/kgu to 1, followed by 44 steps with approximately 1.3 mwd/kgu distance. additional 231 isotopes were added in scale burnup calculation using addnux=3 parameter. serpent used the default isotope inventory setting. 4. results the following parameters were studied: multiplication factor, scattering cross-section (from first group to the second one). in addition to these, fission cross-section, total cross-section and diffusion coefficient, all for two groups. all these parameters were studied for different serpent memory optimization setup. nominal optimization (number 4) was used to produce results which are compared with scale values. 4.1. multiplication factor multiplication factor results are plotted in figure 2a. in order to better show the differences, absolute differences between scale and serpent results were calculated and plotted in figure 2b (scale is the reference one). statistical uncertainty in serpent results is only 22 pcm. uncertainty introduced by nuclear data may reach several hundreds of pcm. however, serpent and scale calculation used nuclear data from an identical library (endf/b-vii.1). scale calculation used preprocessed data in 252 group library, serpent used continuous version. serpent values are higher compared to scale results for boc and eoc. values around 30 mwd/kgu are similar for both codes. initial multiplication factors differ more than 900 pcm from scale results. difference change to 0 160 mwd/kgu is considered as maximal burnup value for currently fuel used. (a). (b). figure 2. multiplication factor from scale and serpent calculation (a) and scale difference from serpent results (b). pcm reached around 30 mwd/kgu. eoc difference reached almost 600 pcm. 4.2. total cross-section total cross-sections for both codes are in good agreement. values are plotted in figure 3a for the fast group and in figure 3b for the thermal group. the serpent code initially calculates the total cross-section by about 0.4 % greater than scale. the value gradually increase during the course of fuel depletion. at burnup level about 20 mwd/kgu the scale value gets greater and the final difference is 0.4 %. results for the thermal group results are almost identical. except for the initial limited period, the differences are limited to 0.1 %. serpent value relative uncertainty is only 0.015% for thermal group and 0.014% for fast group. 4.3. diffusion coefficient the diffusion coefficient is an important macroscopic parameter. diffusion coefficient was calculated for both groups. the coefficient values are plotted in 9 o. novák, j. frýbort, l. sklenka et al. acta polytechnica ctu proceedings (a). (b). figure 3. total cross-section dependency on burnup for fast group (a) and for thermal group (b). figures 4a and 4b. fast group coefficient obtained by scale calculation is higher compared to serpent results. thermal group results are opposite, serpent values are higher compared to scale results. difference for fast group starts approximately at 10% at boc, slowly decreasing and reaching only 3% at eoc. slow group values differs by −7% at boc, drop to −8.6% where it remains till the end. the difference observed in both groups is probably rooted in code structure (stochastic vs deterministic) and in nuclear data difference. serpent value relative uncertainty is only 0.025% for thermal group and 0.021% for fast group. 4.4. fission cross-section fission cross-sections for two energy groups are plotted in figures 5a and 5b. discrepancies from serpent results are plotted in figures 6a (fast group) and 6b (thermal group). fission cross-section for thermal group behaves similarly to the multiplication factor. scale cross-section increases in the first part of fuel cycle, which probably results in decrease of keff difference. maximum is observed similar to minimum (a). (b). figure 4. diffusion coefficient dependency on burnup for fast group (a) and for thermal group (b). of keff difference around 30 mwd/kgu. serpent value relative uncertainty is only 0.038% for thermal group and 0.053% for fast group. 4.5. scattering cross-sections scattering cross-section from fast group to thermal group is plotted in figure 7. both results have similar behavior. serpent values differ from 0.7% to 1% in discrepancy. serpent value relative uncertainty is only 0.014%. 4.6. isotope concentrations number densities of dominant actinides 235u, 238u, and 239pu during fuel burnup were calculated to support understanding of fuel behavior. results are printed in figures 8a, 8b and 8c (including difference from scale results). the observed isotopic changes are consistent but there are differences that can explain differences in the few-group cross-sections. it can be seen that serpent calculates faster 235u depletion than scale and also 239pu number density is lower for the majority of burnup levels. this con10 vol. 19/2018 scale and serpent two-group cross-section data generation (a). (b). figure 5. fission cross-section dependency on burnup for fast group (a) and for thermal group (b). tributes to different and total cross-sections calculated by these two codes. 4.7. memory optimization option in serpent code serpent 2 allows modify calculation in the ram requirements using memory optimization. default mode 4 is recommended for cross-section and other nuclear data generation and has the highest memory requirements. modes 3, 2 and 1 have lower memory requirements and extend calculation duration. more differences are described in serpent forum [1]. in order to compare memory optimization, multiplication factor and other nuclear data were calculated with different optimization. results obtained are similar for all cases and the difference is under statistical uncertainty. the difference for multiplication factors are smaller than 70 pcm which within three standard deviations of the calculation. 5. conclusion monte carlo code serpent and triton calculation sequence from the scale package were used to solve (a). (b). figure 6. scale fission cross-section discrepancy to serpent values for fast group (a) and for thermal group (b). the identical task: calculations of multiplication factor and several two-group macroscopic parameters including scattering cross-section, fission cross-section, total cross-section and diffusion coefficient during depletion of a vvantage-6 fuel assembly. in addition, fuel isotopic composition evolution of selected nuclides was determined. both codes use different methods for solving this neutron transport and depletion coupled task. the results obtained were compared and differences were discussed for each studied parameter. the multiplication factor obtained by both codes are compared in the figure 2a and the differences are plotted in 2b. the initial multiplication factors determined by serpent differs by more than 900 pcm from the scale code results, then it decreased to 0 pcm at around 30 mwd/kgu, however the eoc difference reached almost 600 pcm. total cross-section for both codes is in a good agreement. values are plotted in figure 3a for the fast group and in figure 3b for the thermal group. the relative difference is between ±0.4 %. 11 o. novák, j. frýbort, l. sklenka et al. acta polytechnica ctu proceedings figure 7. scattering cross-section from fast group to thermal group, dependency on burnup. the next studied parameter was the diffusion coefficient. results for the fast group are shown in figure 4a and for the thermal group in 4b. higher differences can be seen for this parameter. they start for the fast group at approximately 10% for boc, slowly decreasing and reaching only 3% at eoc. the thermal group values differs by −7 % at boc, than dropping to −8.6 % where it remains till the end of fuel depletion. the fission cross-section results are plotted in figure 5a for the fast group and in in figure 5b for the thermal group. differences between codes shown in figures 6a and 6b have similar trends and serpent results are consistently lower. the last macroscopic parameter studied was the scattering cross-section. obtained values are shown in figure 7. both codes determined similar burnup dependence with serpent values being different from the scale results by 0.7% to 1%. in addition to macroscopic cross-sections, several isotopic concentration were calculated and studied as well. number densities of dominant actinides 235u, 238u, and 239pu during fuel burnup were calculated. results (including serpent differences from the scale results) can be seen in figures 8a, 8b and 8c. isotopic changes are mostly in agreement, however, serpent calculates faster 235u depletion than scale. moreover, 239pu number density is lower for majority of burnup levels. this contributes to above observed differences of fission cross-section between these two codes. the memory optimization option in serpent code was studied as a part of this study. all four memory optimization were tested. results for all cases were similar and therefore it can be stated that different memory optimization do not affect the results. only the memory requirements and required computational time is affected. in summary, serpent and scale code were compared for burnup calculation and macroscopic crosssection data generation for vvantage-6 fuel assembly. both codes provided mostly similar results. some (a). (b). (c). figure 8. isotope concentration of (a) 235u, (b) 238u and (c) 239pu for both codes and difference between them. differences were however observed in isotopic compositions. it is also reflected in the relevant macroscopic cross-sections. acknowledgements the paper was supported by the project “strengthening and development of research at czech technical uni12 vol. 19/2018 scale and serpent two-group cross-section data generation versity in prague with the use of research infrastructure vr-1 training reactor for research activities” (reg. nr. cz.02.1.010.00.016_0130001790) supported by operational programme research, development and education co-financed from european structural and investment funds and from state budget of the czech republic. support from iaea fellowship number c6/czr/14007 was used in this work. references [1] discussion forum for serpent users [online]. 2007 [accessed on 2017-04-15]. available at: http://ttuki.vtt.fi/serpent/index.php [2] j.r. secker, r.w. miller, l.t. mayhue, r.n. milanova: "advanced design and operating strategies to enhance the safety, operability, and efficiency of vver-1000 reactors." nuclear science and engineering, 121 (1995) 142-145. [3] o. huml aplikace kódů mcnp/x pro výpočet vyhoření a jeho vlivu na neutronově fyzikální charakteristiky reaktorových systémů, neutron physics and burnup calculation of nuclear reactor systems by mcnp/x code . prague, 2014. disertation (ph.d.). czech technical university in prague, faculty of nuclear sciences and physical engineering, department of nuclear reactors (available in czech only). [4] o. novák. study of vver 1000 fuel cycle using nestle and origen codes . prague, 2016. diploma thesis (ing.). czech technical university in prague, faculty of nuclear sciences and physical engineering, department of nuclear reactors. available at: https://www.academia.edu/32855663/study_of_vver_1000_fuel_cycle_using_nestle_and_origen_codes [5] b.t. rearden and m.a. jessee. scale code system, ornl/tm-2005/39, version 6.2.1, oak ridge national laboratory, oak ridge, tennessee (2016). available from radiation safety information computational center as ccc-834. [6] e. fridman, j. leppänen, "on the use of the serpent monte carlo code for few-group cross section generation," annals of nuclear energy, 2011, 38(6), 1399-1405. doi: 10.1016/j.anucene.2011.01.032. issn 0306-4549. [7] m.b. chadwick, m. herman, p. oblozinsky, et al.: "endf/b-vii.1 nuclear data for science and technology: cross sections, covariances, fission product yields and decay data." nuclear data sheets, 2011, 112(12). doi: 10.1016/j.nds.2011.11.002. 2887-2996. [8] t. lötsch, "fuel assembly burnup calculations for vver fuel assemblies with the monte carlo code serpent." kerntechnik, 2014, 79(4), 295-302 [cit. 2017-02-16]. doi: 10.3139/124.110455. issn 0932-3902. [9] j. leppänen, m. pusa, t. viitanen, v. valtavirta, t. kaltiaisenaho: "the serpent monte carlo code: status, development and applications in 2013," ann. nucl. energy, 82 (2015) 142-150. [10] j. leppänen, a. isotalo: "burnup calculation methodology in the serpent 2 monte carlo code." in proc. physor 2012, knoxville, tn, apr. 15-20, 2012. [11] mecir, v. "temelin npp fuel experience". 7 international conference on wwer fuel performance, modelling and experimental support, bulgaria, 2008. 13 acta polytechnica ctu proceedings 19:7–13, 2018 1 introduction 2 fuel assembly description 3 calculation setup 4 results 4.1 multiplication factor 4.2 total cross-section 4.3 diffusion coefficient 4.4 fission cross-section 4.5 scattering cross-sections 4.6 isotope concentrations 4.7 memory optimization option in serpent code 5 conclusion acknowledgements references references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0012 acta polytechnica ctu proceedings 15:12–19, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of drying shrinkage and drying creep kinetics in concrete lenka dohnalová∗, petr havlásek department of mechanics, faculty of civil engineering, czech technical university in prague, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: lenka.dohnalova@fsv.cvut.cz abstract. the aim of this paper is to show and compare the time evolution of drying shrinkage and drying creep in concrete from three different perspectives. the first one analyzes the basic equations defined in the most common design codes and prediction models for the description of the long-term behavior of concrete (aci 209, ec2, model code 2010, b3, b4). next, the evolution of drying creep and shrinkage is examined by processing suitable experimental data available from the database developed at the northwestern university. finally, the last point of view investigates the results obtained from the finite element simulations employing the material point approach, in particular, the material model based on the microprestress-solidification theory. keywords: drying creep, shrinkage, concrete, design code, prediction model. 1. introduction concrete drying is accompanied by its gradual contractive volume changes referred to as shrinkage and additional compliance called drying creep or the pickett effect. on contrary to basic creep (i.e. creep of hygrally sealed specimens at room temperature), both drying shrinkage and drying creep are believed to be bounded. the complex interplay between shrinkage and creep determines the evolution of stresses and so its thorough analysis can help to identify potential cracking and thus can contribute to estimation of the lifetime of the concrete structure. this is the case for members restrained either externally (statically indeterminate structures) or internally (nonuniform stress field generated by the steep gradient of the profile of relative humidity) but very often by the combination of both. this paper summarizes and extends the final thesis of the first author [1]. the first part of this paper compares the time evolution of drying creep and drying shrinkage according to the most common design codes and prediction models. these models operate on the cross-sectional level and thus they provide the estimated average values of strain and compliance which are assumed to be uniform across the cross-section. nowadays, most of the prediction models use as the input parameters the recipe of the concrete mixture, the 28-day compressive strength fc, and the duration of curing (or the onset of drying). in order to estimate the rate of drying (and so the rate of shrinkage and drying creep), it is necessary to provide the model with the size (usually by the quantity called “equivalent thickness” computed as the ratio of volume to drying surface) and the shape of the cross-section. the second part of the paper examines and processes the available experimental data gathered in the database developed and maintained at the northwestern university (nu) [2]. the next part compares the drying shrinkage and drying creep kinetics obtained from the results of the finite element simulations. these simulations operate on the material point level which implicates the non-uniform stress field. if such model is correctly calibrated, it is applicable not only for the analysis of complex and creep-sensitive structures, but also for the assessment of potential surface or internal cracking stemming from the severe drying conditions or excessive temperature generated by cement hydration in massive structures. the results are compared and discussed in the last part of the paper. 2. design codes and prediction models two design codes (aci 209.2r–08 [3] and eurocode 2 [4]) and three prediction models for the long–term behavior of concrete (b3 model [5], b4 model [6] and fib model code 2010 [7]) were selected for a detailed analysis of time evolution of drying creep and drying shrinkage. the investigated prediction models split concrete creep additively into basic creep and drying creep components allowing to evaluate drying creep directly. on the other hand, the analyzed design codes (aci 209, eurocode 2) provide only the formula for the total compliance; therefore, the drying creep compliance was obtained by subtracting the compliance at sealed conditions from the total compliance corresponding to the drying conditions. the total shrinkage is usually considered as the sum of autogenous and drying shrinkage. the b3 model neglects the autogenous shrinkage, for this reason the drying shrinkage is set equal to the total shrinkage. 12 http://dx.doi.org/10.14311/app.2018.15.0012 http://ojs.cvut.cz/ojs/index.php/app vol. 15/2018 comparison of drying shrinkage and drying creep kinetics in concrete the aci 209 does not neglect the autogenous shrinkage but provides only the formula for total shrinkage which is nil at 100 % relative humidity. therefore, similarly to the b3 model, the total shrinkage is taken as the drying shrinkage. even though the individual models use different functions to define the time evolution of drying creep and shrinkage, the shape of the resulting curves is very similar. bigger differences are observed in the final values but these are not in the scope of the present paper. next, we present a brief description of the design and prediction models used in this study. throughout the different models εsh,d stands for the drying shrinkage and henv for the ambient relative humidity (expressed as decimal). the expressions for the drying creep compliance jd are for simplicity presented with t′ = t0, i.e. with the age at loading t′ coinciding with the onset of drying t0. the duration of loading/drying is denoted as t̂ = t− t′ = t− t0. in some models the evolution of drying is derived from v/s which is the ratio of the volume of the concrete member to its surface exposed to the ambient relative humidity. b3 model was developed by bažant and co–workers at the northwestern university and uses, similarly to the other prediction models, empirical formulae to determine the time development of creep and shrinkage. the parameters in the model can be estimated based on the composition of the concrete mixture, the mean compressive strength and in the case of drying on the ambient relative humidity and on the size and shape of the concrete member. the autogenous shrinkage is neglected in the model and so the range of applicability is limited to the water-to-cement ratio above 0.35. the drying creep compliance and drying shrinkage are defined by the following two equations jd(t̂) = q5  e8 ( (1−henv)·tanh √ t̂ τsh −1 ) − 1  0.5 (1) εsh(t̂) = −ε∞sh ·kh · tanh √ t̂ τsh (2) here, τsh is the shrinkage half–time depending on the size and shape of the member, ε∞sh is the absolute value of the ultimate drying shrinkage at henv = 0, kh is a humidity-dependent factor, and q5 is an adjustable parameter. b4 model builds on the previous b3 model and uses more input parameters of the concrete mixture (e.g. aggregate type, additives and admixtures) as well as more elaborate formulae for the prediction of the model parameters. nevertheless, the structure of basic equations of the b4 model is very similar to its predecessor. the autogenous shrinkage and drying shrinkage are treated in the b4 model separately. fib model code 2010 the next two equations present a simplified form of the drying creep compliance and shrinkage functions jd(t̂, t′) = 1 e28 β∗dc [ t̂ βd + t̂ ] √t′ 3.5+ √ 5.29 t′ (3) εsh,d(t̂) = εsh,d0 βrh [ t̂ 0.035 d2 + t̂ ]0.5 (4) where e28 is the modulus of elasticity at the age of 28 days, d is the effective cross section thickness (twice the volume-surface ratio), βd is a size dependence factor, βrh is a humidity dependence factor, and β∗dc and εsh,d0 are model parameters depending on compressive strength, relative humidity and the age at loading. eurocode 2 does not specify an explicit formula for the evaluation of the compliance function j from the creep coefficient ϕ, the conversion was done using a modified expression from model code 1990 [8]. j(t̂, t′) = 1 ecm(t′) + ϕ 1.05ecm(28) [ t̂ βh + t̂ ]0.3 (5) εsh,d(t̂) = kd ·εsh0 · t̂ t̂ + 0.04 ·d1.5 (6) where ecm(t′) is the secant modulus of elasticity of concrete at the age of loading, βh is a parameter depending on relative humidity and member size, kd ·εsh0 expresses the ultimate value of drying shrinkage. aci 209.2r–08 on contrary to ec2, the american standard determines the compliance function from the creep coefficient and the secant modulus of elasticity at the age at loading ecm(t′) j(t̂, t′) = 1 ecm(t′) · [ 1 + t̂ψ d + t̂ψ ·ϕu ] (7) εsh(t̂) = εshu · t̂α f + t̂α (8) where d, ψ, f, and α are parameters, and ϕu and εshu are the ultimate creep coefficient and the ultimate shrinkage, respectively, both influenced by the ambient relative humidity. there are two distinct options how to select the model parameters defining the time evolution of drying creep and shrinkage. the first one, as simplified in [3] from the former version of the code [9], is to use the average and fixed values of the parameters, then ψ = 0.6, α = 1.0 and in general d 6= f. the second possibility, which reflects the size and shape of the structural member, is to set the exponents α = ψ = 1.0, and d = f which are determined from v/s of the drying member. 13 lenka dohnalová, petr havlásek acta polytechnica ctu proceedings basic parameters fc ecm28 t0 t′ c w/c a/c cem. henv 2v/s mpa gpa day day kg m−3 type mm example from aci [3] 33.3 28.2 7 7 409 0.50 4.23 i 0.7 200 experiment troxell [10] 16.5 na 28 28 320 0.59 5.67 i 0.5/0.7 45 experiment bryant [11] 50.1 29.8 8 8 390 0.47 5.09 i 0.6 150/75 ∗) experiment keeton [12] 45 25.9 8 8 452 0.46 3.74 iii 0.2/0.5 65 ∗) 150 mm slab, 75 mm prism table 1. details of the experimental setup in the creep and shrinkage experiments. 10-1 100 101 102 103 104 0 10 20 30 40 10-1 100 101 102 103 104 0 100 200 300 400 500 600 figure 1. drying creep compliance curves (top) and drying shrinkage curves (bottom) given by the analyzed design codes and prediction models. the time evolution of drying creep and shrinkage shown in fig. 1 was obtained with the parameters computed for the concrete mixture and the specimen properties defined in appendix c of aci 209 [3]. the parameters are listed in tab. 1. the specimen is considered as a slab with 200 mm thickness and moist cured for 7 days. the aggregate type used in the b4 model is the quartzite. the curves in fig. 1 are evaluated for the duration of 100 years which is the typical design period of most of the bridges and engineering structures. both figures show the horizontal time axis in the logarithmic scale which is more suitable for processes with initially very high rate and which gradually decreases. fig. 2 presents the time evolution of normalized dry100 102 104 0 0.5 1 100 102 104 0 0.5 1 100 102 104 0 0.5 1 100 102 104 0 0.5 1 figure 2. time evolution of normalized drying creep and shrinkage according to the selected design codes and prediction models. the functions are normalized with respect to the ultimate value. decreasing drying creep from ec2 model is physically unjustified and stems from the model formulation. ing creep, (j̄d, solid red lines) and drying shrinkage (ε̄sh,d, dashed blue lines) for the selected design codes and prediction models. the evolution is normalized with respect to the ultimate value. the presented response of the models is based on the input parameters specified in the first line of tab. 1. naturally, the curves will be different for a modified composition, dimensions or other conditions, but the dominant trends will prevail provided that the ambient relative humidity is kept constant (see the first part in section 5). a different way of comparison is shown in fig. 3 where the normalized drying creep is plotted against normalized drying shrinkage. however, here, the real kinetics of both phenomena is not reflected. it must be noted that in the present example (shown in figs. 1–3), the aci model uses the recommended fixed values of parameters (i.e. the model is size independent). 14 vol. 15/2018 comparison of drying shrinkage and drying creep kinetics in concrete 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 figure 3. comparison of normalized drying shrinkage and normalized drying creep evolution according to the selected design codes and prediction models. normalization is done with respect to the ultimate value. 3. experimental data the freely available experimental database developed at the northwestern university [2] was selected as the source of suitable experimental evidence of drying shrinkage and drying creep. some experimental data comprised in this database were used for the calibration of the models presented in the previous section, and so it can be expected that the processed experimental data will very likely resemble the curves in figs. 2 and 3. the database contains 61,930 values measured on 3,308 specimens (1,439 for creep and 1,869 for shrinkage). these values come from 362 experimental surveys (172 creep, 190 shrinkage). some experimental surveys analyzed more than one concrete composition; in total it was 378 in the case of creep and 652 different compositions in the case of shrinkage. only 68 experimental surveys are relevant to this study because only those used the same recipe of the concrete mixture for both creep and shrinkage specimens. for the present purpose it was necessary to further narrow the selection only to those experiments which tested simultaneously drying shrinkage on unloaded specimens, basic creep, and total creep matter. additionally, the loading in the creep experiments had to be applied simultaneously with the beginning of drying of both loaded and unloaded specimens. the database contains only very few surveys which provide the information on autogenous shrinkage and sometimes it is difficult to distinguish between experiments carried out at sealed conditions and experiments where the specimens are immersed in water or exposed to very high relative humidity and thus exhibit swelling. in order to eliminate this deficiency an additional criteria were set on the value of water-to-cement ratio (minimum value w/c = 0.4) and the compressive strength (maximum value fc = 60 mpa), then the autogenous shrinkage can be treated as negligible compared to the magnitude of the drying shrinkage. only 97 sets from 9 experimental surveys met the aforementioned criteria. unfortunately, it was necessary to further reduce the selection because of incorrect or missing pieces of information, or too short duration of the experiment. the duration was judged as sufficient if the rate of drying creep and shrinkage was minimal at the end of the experiment. this was satisfied only in 6 experimental surveys (with 71 specimens) out of which 3 representative are partially presented in this paper. the value of drying creep was determined as the difference between total and basic creep; when needed (in the case of different times of measurement), the value was computed using linear interpolation between two adjacent measurements. the drying creep and shrinkage was afterwards normalized with respect to the last value attained. the suitable experiments are summarized in the following paragraphs. experiments of troxell [10] adopted in this study are part of an extensive experimental survey focused on the influence of different aspects on the development of creep and shrinkage. the experiment started in 1928 and its duration exceeded 20 years. the data used in this study come from the experimental series #3 which examined the influence of relative humidity; the specimens were cylinders with 102 mm diameter and 356 mm height. the basic properties and the composition of the concrete mixture are specified in tab. 1. the nu database does not contain data for t̂ < 12 days and thus completely misses the initial evolution of creep and shrinkage. two triplets of specimens were suitable for the comparison of drying shrinkage and drying creep kinetics, one for 50 %, the second one for 70 % relative humidity. the results are shown in fig. 4. experiments of bryant [11] was initially focused on study of creep and shrinkage in bridge structures. only specimens with dimensions 150 × 150 × 600 mm were suitable for the present study because only this size was loaded at the age when the drying began which was not the case of the remaining sizes. the specimens should have realistically represented structural members slabs and pillars, to achieve this, one set of specimens had both ends and two opposite lateral sides hygrally sealed (specimens referred to as “slab”) which resulted in unidirectional drying. the other set was sealed only at the ends and thus the specimens were drying evenly from all lateral sides, this set is denoted as “prism”. 15 lenka dohnalová, petr havlásek acta polytechnica ctu proceedings 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 figure 4. comparison of normalized drying shrinkage and normalized drying creep evaluated from the experiments of troxell and bryant. only the points represent the experimental data. at the age of 8 days all specimens were transferred to the environment with 60 % relative humidity. unfortunately, during the experiment the aluminum foil used for sealing has corroded and especially the entirely sealed specimens started unintentionally drying resulting into an increase in creep or shrinkage rate. for this reason the autogenous shrinkage was neglected and the total shrinkage was taken as the drying shrinkage. at the later stage of the experiment, the measurements of the basic creep (spoiled by the sealing leak) partially influenced the determination of drying creep. the processed results are for both groups (slabs and prisms) shown in fig. 4. experiments of keeton [12] the suitable experimental data come from the part of research studying the influence of ambient relative humidity, specimen size, and the load level on creep and shrinkage. the research started in 1957 and lasted for 897 days. the data are not from the nu database but instead from the original paper (and are different). in several cases the increments of basic creep are larger than the increments of total creep. this was the problem especially of smaller specimens with lower level of loading and obviously specimens exposed to higher of relative humidity. additionally, both loaded and unloaded companion specimens were not sealed but were subjected to 100 % relative humidity. owing to the aforementioned reasons the smaller specimens, the specimens with low loading level and specimens drying at 75 % relative humidity are omitted from the present analysis. fig. 5 shows processed results obtained from experiments on cylinders 152 × 457 mm drying at 20 % and 50 % relative humidity and loaded at 30 % and 40 % of 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 figure 5. comparison of normalized drying shrinkage and normalized drying creep evaluated from the experiments of keeton at 20 % and 50 % relative humidity and two levels of loading. f′c is the compressive strength at the age of loading. points are the experimental data. the compressive strength at the age of loading, f′c. in this figure, the higher initial level of shrinkage stems from the first measurement in which the total creep was smaller than the basic creep. 4. data from finite elements simulations except for the prismatic specimens 150×150×600 mm shown in the previous section of this study, the bryant’s experimental survey comprises the information on shrinkage and drying creep measured on geometrically specimens ranging from 100×100×400 mm to 400 × 400 × 1600 mm prepared from a single concrete batch. for this reason this experiment was in [13] and [14] selected as an ideal opportunity for verification of the material model based on microprestresssolidification theory (mps) [15]. it turned out that the original mps gives the opposite size effect on drying creep than the experiment and for this reason the mps model was modified [13, 14]. here we present the comparison of drying creep and shrinkage for both approaches. hereafter, the modified version of mps model is referred to as “mps improved”. the presented data use the processed results obtained from the finite element simulations carried out in the open source fe package [16]. the problem was analyzed as one-way coupled with the structural problem depending on the moisture transport according to the bažand and najjar model [17] for diffusion of water vapor. in order to obtain the shrinkage and drying creep kinetics, it was necessary perform, similarly to the real experiment, 3 analy16 vol. 15/2018 comparison of drying shrinkage and drying creep kinetics in concrete 100 102 104 0 0.5 1 100 102 104 0 0.5 1 100 102 104 0 0.5 1 100 102 104 0 0.5 1 figure 6. the processed results of fe simulations of bryant’s experiment showing the time evolution of normalized drying creep and shrinkage. the top two figures are for 150 mm slabs bottom two figures are for 150 × 150 mm prism. the curves are normalized with respect to the ultimate value. ses: basic creep, drying creep and shrinkage. the values of the material parameters in the models were selected in order to obtain a reasonably accurate fit of all specimens (“prisms” and “slabs” of different sizes), therefore, the agreement of the simulation of 150 mm specimen with the experimental data cannot be expected to be satisfactory. the normalized drying creep and shrinkage is for the original and improved mps model and the two geometries shown in fig. 6. 5. discussion in general, according to the prediction models and design codes, the time evolution of drying creep and drying shrinkage depends dominantly on the ambient relative humidity and on the size of the concrete member, the influence of other factors is not so significant. when evaluating the differences in kinetics of drying creep and drying shrinkage, each model is sensitive only to certain input parameters. in the case of the b3 model, the only factor which affects the relationship between drying creep and drying shrinkage is the ambient relative humidity. at low relative humidity, the drying creep is delayed behind shrinkage, but on contrary to this, at high levels (henv > 0.8), the delay is in shrinkage. however, the difference is in both cases relatively small, see fig. 7. on the other hand, the fib mc2010 is completely insensitive to henv. the kinetics is influenced dominantly by the size of the member (its v/s), see fig. 8. the larger the size the more is the curve in fig. 2 bent downwards and the bigger is the delay of shrinkage behind drying creep. this is caused by the fact that with decreasing v/s the drying shrinkage becomes 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 figure 7. the influence of different levels of ambient relative humidity on the drying creep vs. drying shrinkage kinetics according to the b3 model. the remaining parameters are from the first line in tab. 1. 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 figure 8. the influence of different v /s ratios on the drying creep vs. drying shrinkage kinetics according to the fib mc 2010 model. the remaining parameters are from the first line in tab. 1 more accelerated than the drying creep. to certain extent the kinetics is also influenced by the onset of drying, but for reasonable values the differences are not substantial. the eurocode 2 predicts, except for extreme combinations of parameters, delay in shrinkage behind drying creep. the model is sensitive both to v/s a and henv. for smaller specimens and higher relative humidity the drying creep curve becomes more non-monotonic. 17 lenka dohnalová, petr havlásek acta polytechnica ctu proceedings 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 figure 9. the comparison of the experimental data, prediction models and fe simulations of the bryant’s experiment, 2v /s = 150 mm (slab). aci 209 is the exception to the other presented models. for first approach with the average recommended values (according to [9]) the kinetics of drying creep and shrinkage remains constant, with an initial small delay of shrinkage behind drying creep. according to the second approach, which sets equal the exponents α and ψ as well as the parameters d and f, the kinetics of drying creep and drying shrinkage become identical (i.e. the function would be a diagonal in fig. 3). as presented in figs. 4 and 5, the selected suitable experimental data give the opposite trend compared the prediction models (except for b3/4 at lower relative humidity and aci at the later stage of drying). this means that in the experiments the drying shrinkage is a little bit ahead of the drying creep. the points in these figures correspond to the experimental data which are interpolated by the curves of the same color. two issues have to be pointed out. first, the resulting shape of the curve in the normalized plot shrinkage vs. drying creep depends on the duration of the experiment. insufficiently long experiments can lead to misleading conclusions. second, as demonstrated in fig. 5, the readings should be documented more frequently in the initial phase of drying and the specimens should be instrumented as soon as possible. in the case of keeton’s experiment, the first displayed data point corresponds to the drying duration 7 days and the shrinkage strain reaches almost 40 % of the last measured value (after approx. 3 years). (the first data points for t̂ = 1 day which would have yielded negative jd is omitted here.) according to the computations with the original version of the material model based on the microprestresssolidification theory, the drying creep evolves in an 0 0.2 0.4 0.6 0.8 1 0 0.2 0.4 0.6 0.8 1 figure 10. the comparison of the experimental data, prediction models and fe simulations of the bryant’s experiment, 2v /s = 75 mm (prism). unexpected fashion. first, as shown in the two left graphs in fig. 6 and in figs. 9 and 10, the drying creep evolves initially faster than drying shrinkage, then, suddenly, the creep stops and after certain time it continues with a considerable delay after shrinkage. this effect is more pronounced in the case of prismatic shape which dries more rapidly. the response of the improved mps model is more promising, see the two right graphs in fig.6. similarly to the cross-sectional approaches, it gives a small initial delay of shrinkage after drying creep, afterwards, the relationship is almost linear. however, the initial delay is reversed compared to the analyzed experimental data. very probably, in the case of the models working on the cross-sectional approach, the differences of the shrinkage and drying creep kinetics origin from their optimization and calibration procedure. very likely the models were calibrated against the entire database with evenly distributed weights and without emphasizing that some specimens are companions belonging to the same experimental survey. 6. conclusions the presented study analyzed the time evolution of drying shrinkage εsh,d and drying creep jd from three different perspectives: the most common design and prediction models, experimental data measured on laboratory concrete specimens, and the finite element simulations. the following conclusions can be drawn: • even though the design codes and prediction models are calibrated on a very similar experimental database, jd and εsh,d are described in each model 18 vol. 15/2018 comparison of drying shrinkage and drying creep kinetics in concrete differently and their kinetics is different. additionally, in each model, the relationship between the jd and εsh,d depends on different factors. three of the analyzed models, b3, b4, and aci 209, give for the standard conditions comparable kinetics of these processes. however, in the case of fib mc2010 and ec2 the delay of εsh,d behind jd is substantial. according to ec2, jd is a non-monotonic function. • a closer inspection of the experimental database which contains a vast amount of data revealed that only very few experiments fulfill the needs of the present research. unfortunately, most of the experiments does not last sufficiently long or certain essential pieces of information are missing. the analyzed data indicate that the jd is initially slightly delayed behind εsh,d, which means that the trend is exactly opposite compared to the prediction models. • the improved mps model gives more realistic kinetics of jd and εsh,d compared to its predecessor which unexpectedly gives a significant decrease in drying creep rate in the middle of drying. both models predict, similarly to the cross-sectional models, an initial small delay of εsh,d behind jd. • sufficiently long experiments on normal-strength concrete are vital for the development and proper calibration of the material models for creep and concrete, however, they are very scarce. acknowledgements the authors gratefully acknowledge financial support from the grant agency of the czech technical university in prague, project number sgs18/037/ohk1/1t/11. the numerical analyses have been performed with oofem, an open-source object-oriented finite element program [16, 18]. references [1] l. dohnalová. comparison of shrinkage and drying creep kinetics of concrete, bc. thesis. czech technical university in prague, 2018. 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[6] rilem technical committee tc-242-mdc (z.p. bažant, chair). model b4 for creep, drying shrinkage and autogenous shrinkage of normal and high-strength concretes with multi-decade applicability (rilem technical committee tc-242-mdc multi-decade creep and shrinkage of concrete: material model and structural analysis). materials and structures 48:753–770, 2015. [7] fédération internationale du béton. model code 2010. no. vol. 65 in fib bulletin. international federation for structural concrete (fib), 2012. [8] comité euro–international du béton (ceb). ceb–fip model code 1990: design code. t. telford, london, 1993. [9] a. c. 209. prediction of creep, shrinkage, and temperature effects in concrete structures. american concrete institute, 1997. [10] g. troxel, j. raphael, r. davis. long-time creep and shrinkage tests of plain and reinforced concrete. proc astm 58 pp. 1101–1120, 1958. [11] a. h. bryant, c. vadhanavikkit. creep, shrinkage-size, and age at loading effects. aci materials journal 84:117–123, 1987. [12] keeton. study of creep in concrete, phase 1 (i-beam). u. s. naval civil engineering laboratory, port hueneme, california, 1965. [13] p. havlásek. creep and shrinkage of concrete subjected to variable environmental conditions, phd. thesis. czech technical university in prague, 2014. [14] z. bažant, p. havlásek, m. jirásek. microprestress-solidification theory: modeling of size effect on drying creep. in n. bicanic, h. mang, g. meschke, r. de borst (eds.), computational modelling of concrete structures, pp. 749–758. crc press/balkema, eh leiden, the netherlands, 2014. [15] z. p. bažant, a. p. hauggaard, s. baweja, f. j. ulm. microprestress solidification theory for concrete creep. i: aging and drying effects. journal of engineering mechanics 123:1188–1194, 1997. [16] b. patzák. oofem home page, 2000. http://www.oofem.org. [17] z. p. bažant, l. j. najjar. nonlinear water diffusion in nonsaturated concrete. materials and structures 5:3–20, 1972. doi:10.1007/bf02479073. [18] b. patzák. oofem an object-oriented simulation tool for advanced modeling of materials and structures. acta polytechnica 52(6):59–66, 2012. 19 http://dx.doi.org/10.1007/bf02479073 acta polytechnica ctu proceedings 15:12–19, 2018 1 introduction 2 design codes and prediction models 3 experimental data 4 data from finite elements simulations 5 discussion 6 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.9.0026 acta polytechnica ctu proceedings 9:26–31, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app removal of diquaternary ammonium cations from as-synthesized ssz-16 zeolite tatana supinkovaa, c, ∗, ivan jirkaa, jan drahokoupilb, d, jan langmaiera, vlastimil filac, libor brabeca, milan kocirika a j. heyrovsky institute of physical chemistry,czech academy of sciences, dolejskova 2155/3, prague 8, czech republic b faculty of nuclear sciences and physical engineering, czech technical university, trojanova 13, prague 2, czech republic c faculty of inorganic technologies, university of chemistry and technology in prague, technicka 5, prague 6, czech republic d institute of physics, czech academy of sciences, na slovance 1999/2, prague 8, czech republic ∗ corresponding author: tatana.supinkova@jh-inst.cas.cz abstract. zeolites are stable microporous aluminosilicates with numerous applications in chemical technology such as separation of species and catalytic transformations. our study is focused on a weakly explored zeolite ssz-16 with pore constrictions defined by 8-membered oxygen rings. key results are the preparation of et6-diquat-5 dication used as a structure directing agent (sda) and finding the optimum synthesis conditions with respect to zeolite phase purity. stability of sda was examined in conditions similar to those of autoclave synthesis (concentration, ph, temperature). moreover, the content and location of sda species in zeolite phase and conditions of sda decomposition were investigated. keywords: zeolite, hydrothermal synthesis, ssz-16, afx, et6-diquat-52+. 1. introduction zeolites and zeolite-like materials consist of tetraedral building units to4 where t are central atoms most frequently si and al+. tetraedra are linked to each other through oxygen atoms. the spaceous arrangement of tetrahedra form as a rule microporous structure where micropores take the shape of regular voids (channels and/or cavities) which are repeated through the crystals with the regularity of the crystalline structure. in zeolites, t-atoms can be replaced by isomorphous substitution, for example ge, ti, p, fe, ga, b and thus form zeolite analogues. zeolites and their analogues are crystalline materials with regular microporous structure (with pore width < 2 nm). the intracrystalline pores are thus of molecular dimensions. the surface of intracrystalline pores is formed by oxygen atoms. the critical pore width is for atoms and molecules as a rule determined by oxygen rings containing 8, 10 or 12 oxygen atoms. small cations can pass even through 6-membered oxygen rings. micropores can accommodate spherical molecules typically with the diameter of 0.3-1.0 nm or cylindrical molecules with cylinder diameter in the above range. zeolites are also called molecular sieves due to possible exclusion of molecules with size exceeding some critical value. zeolites are perspective materials for industrial applications due to their molecular sieve effect, high adsorption capacity, and frequently to the presence of catalytic sites or to their ion exchange capability. a relatively novel zeolite ssz-16 belongs to the class of small pore zeolites characterized by 8membered oxygen rings which may exhibit molecular sieve effect for gases with small molecules. the main field of its application is believed to be in gas separation of various hydrocarbons and mixtures containing molecules such as co2, h2, ch2 and other gases [1–3]. furthermore, zeolite ssz-16 can be used as microporous layers and particles to develop composite and mixed matrix materials [4–6]. zeolite ssz-16 was also tested as a catalyst. for example, cu-ssz-16 [7, 8] seems to be promising as a selective catalytic reduction catalyst for the direct conversion of no into n2 and o2. zeolite h-ssz-16 was found to be highly selective to dimethylamine in the catalytic reaction of ammonia and methanol [9]. the principal aim of this work is to optimize zeolite ssz-16 hydrothermal synthesis and contribute to answering questions concerning limits to phase purity, yield and posibility to control morphology and size of crystals. moreover, the research could clarify the role of diquaternary ammonium species in the synthesis. the principal aim of this work is to contribute to the knowledge of structure-directing agent (diquaternary ammonium dications) chemistry in relation to its (i) synthesis and purity, (ii) role in the synthesis of ssz-16 zeolite, (iii) stability in synthesis batches and (iv) its thermal removal from the zeolite framework. 26 http://dx.doi.org/10.14311/app.2017.9.0026 http://ojs.cvut.cz/ojs/index.php/app vol. 9/2017 removal of diquaternary ammonium cations from as-synthesized ssz-16 zeolite figure 1. left afx framework type representation using cbu and their arrangement (adapted from www.iza-online.org), right afx framework type representation using 6-6 sbu (adapted from ilyushin et al. 2015). figure 2. view of the afx structure along the caxis; yellow spheres t-atoms, red spheres oxygen atoms, blue patterns visualization of electron density distribution. 2. theory 2.1. zeolite ssz-16 and its structure zeolite ssz-16 is isotypic with its zeolite analogue sapo-56 [7, 10, 11]. both these microporous materials exhibit the afx framework type. this framework shows the hexagonal symmetry, the corresponding space group of symmetry is p63/mmc and the unit cell idealized parameters are a = b = 13.7 å, c = 19.7 å. fundamental articles on ssz-16 present the afx framework type in the terms of composite building units (cbu): gmelinite (gme) and aft cavities and d6r (double six-ring) cages [10, 12, 13], see figure 1 (left). this method of afx framework type representation is widely used to visualize spaces capable of accommodating sorbing species. however, this method is less suitable for the analysis of various zeolite framework topology features due to the existence of common t-atoms for the neighbouring cbu. an alternative way of afx framework representation is by secondary building units (sbu) using either an array of 6-membered rings (6 code) or double 6-membered rings (6-6 code), see figure 1 (right). figure 2 represents the view of the afx structure along the c-axis using the visualization of electron density distribution (computed by dft calculation implemented in castep program). zeolites ssz-16 and sapo-56 belong to small-pore zeolites with 3d pore system. the narrowest passages through these pores (also called bottlenecks or configure 3. 8-membered oxygen ring in sapo-56 viewed normal to [001], red spheres oxygen atoms, yellow spheres t-atoms figure 4. structure of sda hexaethylpentane diammonium dication (blue spheres nitrogen n+, grey carbon, white hydrogen). strictions) are formed by 8-membered oxygen rings. dimensions of these rings were published for sapo56 [11], see figure 3. 2.2. role of sda species in synthesis of ssz-16 essential components for zeolite synthesis are silica, alumina, source of alkali, sda and water. theoretical approaches offer a great number of hypothetical zeolite structures. however, only a small fraction of them can be synthesized. besides, a considerable number of zeolite phases can be formed only in the presence of particular organic species termed as structure-directing agents (sda). a mechanism of sda action in synthesis mixtures is still an open question. attempts were made to explain sda effect on the basis of sda species size and shape, its rigidity, hydrophobicity/hydrophilicity and sda species interactions with solution components. diquaternary ammonium species were proposed in the nineties as a novel group of sda species for the synthesis of zeolites. the research in our laboratory showed that hexaethylpentane diammonium dication (c2h5)3n+(ch2)5n+(c2h5)3 (also called as et6-diquat-52+), appears to be, among diquaternary ammonium species, the most efficient sda for the synthesis of zeolite ssz-16 [14]. the model of this sda dication is shown in figure 4. 27 t. supinkova, i. jirka, j. drahokoupil et al. acta polytechnica ctu proceedings figure 5. computer simulated location of sda molecules in uc of ssz-16 a theoretical picture of sda species in ssz-16 structure was derived using the following assumptions: • a non-decomposed sda dication can be accommodated only in the aft cavity and thus oriented with its maximum dimension along the c-axis (see figure 5). • any aft cavity accomodates just a single sda dication. • there are two aft cavities per uc. therefore, the uc contains two sda dications non-decomposed during the synthesis procedure. • in principle, the gme cavities can accomodate only some products of sda decomposition; this phenomenon is not considered in the theoretical estimate of sda content. based on these assumptions, the following formula for sda dication mass fraction (gsda)th in molar mass of uc for non-calcined zeolite ssz-16 can be formulated as (gsda)th = nsda msda (mu c )0 + nsda msda , where nsda stands for the number of sda per uc and is equal to 2, msda denotes the molar mass of sda and equals to 272.32 g mol−1. (mu c)0 represents the molar mass of zeolite ssz-16 free of sda and can be expressed depending on the atomic ratio si/al = z as (mu c )0 = 48 mn a 1 + z + 48 msi z 1 + z + 48 mal 1 + z + 96 mo, for z = 6 and (mu c)0 = 3029.23 g mol−1. based on these equations, the sda dication mass fraction (gsda)th can be calculated and it equals to 0.152. the experimentally measured sda content (gsda)th in the zeolite synthesis product will be compared to this value. 3. experimental 3.1. synthesis of sda in view of the fact that the price of species containing et6-diquat-52+ dication appeared to be prohibitive, a protocol for its synthesis was developed. the reactants used for this purpose were 1,5-dibromopentane and triethylamine. the reaction proceeded in ethanol solvent. the crystallization was performed in diethyl ether solvent. because the synthesis proceeds in twostages, the purity of the synthesis product were examined by x-ray diffraction and by thermogravimetry. the heating program for sda thermogravimetry was the same as for zeolite samples calcination and zeolite thermogravimetry, see below. sda with significant content of the first stage synthesis product can thus be rejected or recycled. 3.2. synthesis of ssz-16 particles the research was directed to the development of a static in-situ hydrothermal crystallization procedure of ssz-16 zeolite [14, 15]. the synthesis was performed in three steps: preparation of starting synthesis mixture, ageing of synthesis mixture (formation of viable nuclei), and crystallization of aged mixture. the most suitable silica source appeared to be colloidal silica. the starting synthesis mixture was composed of colloidal silica ludox as 30 (30% sigma-aldrich), aluminum nitrate nonahydrate, sodium hydroxide, organic template et6-diquat-5 br2 and deionized water [14, 16]. the synthesis process was adapted by changing the synthesis solution composition, ageing period, temperature and duration of crystallization [15, 16]. the composition of synthesis solution was varied and optimized to obtain pure phase of ssz-16 particles. the ageing period was performed for seven days to promote nucleation of precursor particles (formation of viable nuclei). the ageing process of ssz-16 was performed at elevated temperature with the synthesis solutions heated up to 80 °c inside an oil bath, continuous stirring was ensured by a magnetic stirrer inside teflon vessels. the crystallization process was performed for seven days, temperature adjusted to 160 °c. the hydrothermal in-situ synthesis of ssz-16 crystalline particles performed inside teflon-lined stainless steel stationary autoclaves from the pre-aged solutions under autogenous pressure. the crystalline ssz-16 products were purified inside an ultrasonic bath, washed and dried in a pre-heated oven overnight. after the synthesis, the sda was removed by the two-cycle thermal calcination process under nitrogen/air atmosphere to open pores inside the aft cage. the temperature program for both the cycles was as follows: heating with the rate 0.5 °c/min to 120 °c, keeping at 120 °c for two hours, heating to the temperature 550 °c with the rate 0.5 °c/min, then keeping at 550 °c for 8 hours and cooling to the room temperature with the rate 0.5 °c/min. 28 vol. 9/2017 removal of diquaternary ammonium cations from as-synthesized ssz-16 zeolite 3.3. characterization zeolite ssz-16 particles were characterized by scanning electron microscopy (sem) technique using jeol jsm 5500lv to examine crystal morphology and particle size. powder x-ray diffraction (xrd) technique was used to determine crystallinity and phase purity of zeolite ssz-16 particles. x-ray diffraction was performed on panalytical x’pert pro diffractometer with co anode (wavelength λ = 0.1789 nm) in bragg-brentano geometry. the fixed divergent slits in primary beam and 1d detector x’celerator in diffracted beam were used. the spectra were collected in the 2θ range of 5-30°. the measurements and analyses of collected spectra were performed in the institute of physics, cas. the micropore volume of calcined zeolite ssz-16 particles was evaluated from nitrogen adsorption-desorption isotherms. the isotherms were measured on asap 2020 (micromeritics, usa) volumetric instrument at -196 °c. the sample was degassed under a vacuum at 350 °c for 8 h prior to the analysis. the micropore volume was evaluated by the t-plot method. particular attention was paid to sda. it concerned sda purity (xrd analysis), thermal stability and its decomposition both in synthesis solutions and products. thermal stability of sda in alkaline solutions were analyzed after the heat treatment in an autoclave. sda concentration was measured using cyclic voltammetry (cv) [17] with the home-made apparatus in j. heyrovsky institute of physical chemistry, cas. sda content in zeolite samples and sda residua after their calcination were measured using elemental analysis performed on elementar vario el iii in chns operation mode, and further on with x-ray photoelectron spectroscopy xps using esca 3 mk ii spectrometer (vg) in fixed transmission mode. the kinetics of sda removal/decomposition and related heat effects were investigated with thermogravimetry and differential scanning calorimetry using setsys evolution tga, setaram. 4. results and discussion among silica sources for synthesis mixtures, ludoxbased synthesis mixtures had the best properties both from the point of view of hydrogel density and capacity of crystallization. optimum synthesis batch composition to eliminate unwanted phases in product was determined as 3 sda: 15 na2o: 0.5 al2o3: 30 sio2: 1200 h2o. deviation from this composition causes considerable changes in phase composition and formation of different zeolitic phases. the introduction of ludox as 30 as silica source had two consequences: (i) reducing the size of amorphous core of zeolite particles which was observed in the early stages of synthesis procedure development (see figure 6a) and (ii) dramatic reducing the size of particles. the size of particles decreased to approximately 2 µm and well-developed polycrystalline particles formed (see figure 6b). this favourable development of product quality manifests itself also (i) in x-ray spectra (see figure 6. a amorphous core and crystalline shell of a particle (2 000×), b well developed polycrystalline particles (10 000×). figure 7. xrd spectra of selected samples and theoretical spectrum of ssz-16 figure 7) and (ii) in sda content measured gravimetrically from mass difference between non-calcined and calcined samples, showing on selected zeolite samples the development of (gsda)exp. in contrast to earlier preparations where amorphous cores were visible on disrupted particles and the values of (gsda)exp ranged between 0.160 and 0.180, in more recent experiments the amorphous cores were not observed and (gsda)exp ranged between between 0.140 and 0.150 approximately. it is suspected that amorphous cores exhibited a higher content of sda as compared with zeolite phase. figure 7 represents an example of the x-ray diffraction spectra of selected calcined samples compared to the theoretical diffraction spectrum for the zeolite ssz-16. the samples were confirmed to be crystalline and pure phase ssz-16 particles. regarding the content of organic species in the synthesized zeolite, a question arises whether some decomposition products of sda cannot be located in gme cavities. the evaluation of sda thermal stability in alkaline solutions performed by cyclic voltammetry showed a relatively rapid sda decomposition. the time dependence of sda concetration during its thermal treatment in alkaline solution at temperature, ph, and sda concentration close to those in synthesis mixtures is shown in figure 8. it is obvious that during the reaction time 6 to 8 hours the sda concentration decreases to a value lower than 50 % of its initial value. this may generate a considerable number of sda fragments and enhance the mass of organics incorporated in ssz-16 crystals. nevertheless, for optimized synthesis procedure the values of (gsda)exp remain in the range 0.140 to 0.150 which cast doubt on additional places of organics accommodation in ssz-16 crystals. accessibility of microporous crystal cavities after 29 t. supinkova, i. jirka, j. drahokoupil et al. acta polytechnica ctu proceedings figure 8. kinetics of et6-diquat-52+ decomposition at 160 °c. initial concentration of et6diquat-5 br2 c0sda = 0.168 mol dm−3, initial concentration of naoh c0sda = 1.66 mol dm−3, solvent deionized water. figure 9. n2 adsorption-desorption isotherm. full circles adsorption, empty circles desorption. template removal was confirmed by measurements of adsorption and desorption isotherms of n2 at -196 °c. this result is exemplified by isotherms measured on calcined sample (figure 9). the isotherm shows a sharp rise in n2 adsorption at low relative pressure and negligible hysteresis loop. its shape shows that the material is microporous with negligible effect of mesopores. the sample showed the micropore volume of 0.24 cm3 g−1. at the same time, the elemental analysis excluded presence of organic residua in crystal bulk after ssz-16 calcination. this is a contrast to the situation in crystal subsurface region where carbonaceous deposits were detected after zeolite calcination by xps. the thermogravimetric curves monitored during sda removal from zeolite samples using the above temperature program are exemplified on figure 10. temperature calibration was performed using cuso4.5h2o and sn. on the picture is the comparfigure 10. examples of tg curves percentage of the sample mass based on initial mass vs. temperature; red silicalite-1 templated with tpa+; blue ssz-16 templated with et6-diquat-52+. ison of sda removal from silicalite-1 (tpa+) with that from ssz-16 (et6-diquat-52+). there is bigger amount of organics in ssz-16 and steeper decay in the first portion of the tg curve. we suspect that it may be related to more rapid thermal decomposition of et6-diquat-52+ in synthesis batch as compared with tpa+ in silicalite-1 which is considerably more stable. we also suspect that sda fragments may be incorporated into gme cavities. 5. conclusions the principal contribution of this work is the feasibility of zeolite ssz-16 synthesis. the selected sda hexaethylpentane diammonium cation turned out to be appropriate and the synthesis route to this species was elaborated. by variation of synthesis parameters of aging and crystallization, the particle size was reduced from cca 50 µm to approximately 2 µm. moreover, the conditions to attain purity of ssz-16 higher than 98 % were found and polycrystalline samples with well developed morphology were obtained. using n2 adsorption and desorption at -196 °c, volume of micropores vmicro = 0.24 cm3 g−1 was evaluated. an insight into the chemistry of sda was presented. in particular, in its synthesis and conditions of its purity. further on, a novel technique was introduced to measure concentration variation of cationic sdas under conditions close to those at synthesis conditions. particular attention was also paid to the investigation of template removal kinetics. acknowledgements the financial support by czech science foundation via grant ga16-02681s is gratefully acknowledged. references [1] m. pera-titus et al. porous inorganic membranes for co2 capture: present and prospects. chem rev 114(2):1413–1492, 2014. doi:10.1021/cr400237k. 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[17] j. langmaier et al. extreme basicity of biguanide drugs in aqueous solutions: ion transfer voltammetry and dft calculations. j phys chem a 120:7334–7350, 2016. doi:10.1021/acs.jpca.6b04786. 31 http://dx.doi.org/10.1021/cm4015095 http://dx.doi.org/10.1016/j.seppur.2007.09.008 http://dx.doi.org/10.1039/c4ta02744b http://dx.doi.org/10.1016/j.apcatb.2010.12.022 http://dx.doi.org/10.1039/c4cc09645b http://dx.doi.org/10.1016/j.apcata.2006.02.044 http://dx.doi.org/10.1021/cm960289c http://dx.doi.org/10.1016/s1387-1811(98)00293-5 http://dx.doi.org/10.1021/cs300558x http://dx.doi.org/10.1016/s1387-1811(01)00441-3 http://dx.doi.org/10.1016/s1387-1811(03)00381-0 http://dx.doi.org/10.1016/j.micromeso.2016.03.033 http://dx.doi.org/10.1016/j.micromeso.2013.03.007 http://dx.doi.org/10.1021/acs.jpca.6b04786 acta polytechnica ctu proceedings 9:26–31, 2017 1 introduction 2 theory 2.1 zeolite ssz-16 and its structure 2.2 role of sda species in synthesis of ssz-16 3 experimental 3.1 synthesis of sda 3.2 synthesis of ssz-16 particles 3.3 characterization 4 results and discussion 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.14.0014 acta polytechnica ctu proceedings 14:14–20, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app calculation uncertainties in spent fuel inventory determination tomas czakoja, ∗, jan fryborta, martin loveckyb a czech technical university in prague, fnspe, dnr, v holesovickach 2, 18000 prague, czech republic b skoda js a.s., orlik 266/15, 31600 bolevec, plzen, czech republic ∗ corresponding author: tomas.czakoj@gmail.com abstract. in the depletion calculation of the nuclear fuel, the uncertainty is of utmost importance, as it affects the uncertainty of the subsequent calculation, when the calculated composition is used. the calculations are even more important when they are safety related, e. g., when determining the reactivity or emissions of radioactivity to the environment. this work covers the depletion model of gd-2m+ fuel which was developed in origen-arp/triton calculation sequences, both being parts of a scale 6.2.1 package. the uncertainties of the respective calculation model were determined by comparison with experimental data for both calculation sequences. the effects of operational and manufacturing uncertainties on concentration of the most important nuclides using triton depletion model of gd-2m+ fuel were determined. the effect of respective uncertainties manifested in changes of composition on the multiplication factor using monte-carlo sequence keno-vi were also specified. keywords: uncertainties, nuclear, fuel, calculation, depletion. 1. introduction the awareness of the uncertainty of a calculation is very important, especially in nuclear reactor applications. inaccurate determination of spent fuel inventory can result in higher decay power, higher emissions or by increased reactivity which can be very dangerous. this work estimates the accuracy of the calculation performed by comparison of experimentally determined spent nuclear fuel inventory with the calculated one. the study of influence of variation in some operational and manufacturing input aspects, such as temperature, power level or enrichment, was conducted. the respective study focuses on the spent fuel inventory determination and calculation of its multiplication factor. 2. calculation methods all calculations described in this paper were made in scale code system version 6.2.1 [1]. the depletion was simulated in triton and origen-arp sequence, the determination of the reactivity important nuclides was carried out in tsunami-2d sequence and the determination of the multiplication factor was obtained with the use of the keno-vi stochastic sequence. both depletion and sensitivity calculations used a multigroup version of endf/b vii.1 nuclear data library, keno-vi simulation used continuous energy version of the aforementioned library [1, 2]. 2.1. depletion model two depletion models of gd-2m+ fuel were created. the fuel used in the depletion models is currently used in vver-440 reactors in dukovany power plant. this fuel assembly contains fuel pins without a burnable figure 1. enrichment of each fuel pin in the gd2m+ fuel assembly. numbers mean enrichment in weight fractions and the hexagon with brown color shows position of fuel pin with burnable absorber. [3]. absorber which does not have a central hole in the fuel pellets. only exception is six fuel pins with gd2o3 which contain a central hole [3]. figure 1 shows the enrichment of each fuel pin in the respective fuel assembly, as well as the position of central tube (ct) and fuel pin with burnable absorber gd2o3. the numbers in figure 1 mean enrichment in weight fractions and the hexagon with brown color shows position of fuel pin with the burnable absorber [3]. the first depletion model was created in triton sequence. the detailed model included the description of geometry, composition, power history and several other important parameters affecting the accuracy of 14 http://dx.doi.org/10.14311/app.2018.14.0014 http://ojs.cvut.cz/ojs/index.php/app vol. 14/2018 calculation uncertainties in spent fuel inventory determination the calculation. this model contained an infinite array of the fuel assemblies with b1 buckling, where the assumed spectrum is identical to the critical one [4]. the depletion model used five cycles with a duration of 330 days with the specific power set to 30.535 w g−1hm and the shutdown between cycles of 35 days. the time step between depletion was 1.2 days for first 6 days of cycle, next 24 days used 6 days long time step and following 300 days used 20 days long time step. the depletion model was originally set to maximum accuracy and after the creation of the model, many optimization calculations were conducted. each calculation used the basic model with one modified parameter thus quantified the effect of selected reduction. in case the effect was negligible, the reduction was implemented in the final model. the optimized final model was twice faster than the basic one and the accuracy of the results was not affected [1]. the optimized parameters were time steps or legendre order in the transport theory. next optimized parameter was addnux, which adds trace quantities of certain groups of nuclides to the inventories of depletion materials which causes better tracking of these nuclides during the calculation. last optimized parameter was assign. this parameter groups depletion materials. each material of the respective group is independently depleted, but the calculation uses average cross section. the second depletion model was made in origenarp sequence. this sequence, together with the model, is very fast and simple, however it requires a data file from the previous triton calculation. the results are similar to results from slow detailed triton calculation, if an appropriate data file is used. the origen-arp model contained only a homogeneous composition and the power history, no other parameters are required [1]. 2.2. comparison with experimental data because there were no avaliable experimental data of the spent gd-2m+ fuel, it was nessesary to use the data of another fuel type. the data used for the comparison were retrieved from project istc # 3958. the goal of the project was to obtain data from the spent nuclear fuel for description of its composition [5]. the fuel assembly used in the istc project differed from the gd-2m+ fuel assembly in several aspects. the first difference was in enrichment because istc fuel assembly contained fuel pins with constant enrichment. next difference was in the geometry because istc fuel contained central hole and also had moderately different dimensions. because of these distinctions, the istc fuel assembly contained smaller amount of uranium than the gd-2m+ fuel assembly. another significant difference was the absence of any burnable absorber in istc fuel. all these diferences are listed in table 1 [3, 5, 6]. gd-2m+ istc #3958 fuel enrichment 3.6/4.0/4.6 % 4.4 % outer fuel diameter 0.78/0.76 cm 0.757 cm inner fuel diameter 0/0.12 cm 0.12 cm inner cladding radius 0.793/0.773 cm 0.772 cm burnable absorber gd2o3 table 1. differences between gd-2m+ and istc fuel 157gd 156gd 158gd 159tb 155gd 161dy 154gd 239pu 238u 241pu 240pu 135xe 103rh 143nd 133cs 237np 149sm 131xe 242pu 236u 99tc 155eu 153eu 151sm 154eu 152sm 147pm 145nd 243am 109ag 101ru 95mo 238pu 148mpm 150sm 105pd 147sm 134cs 105rh 141pr 139la 241am 235u 234u 16o 242cm* 244cm* 108pd* table 2. nuclides with big impact on the reactivity the istc project # 3958 published information about 12 samples which came from the above mentioned fuel assembly. it includes information about burnup, location of samples in the fuel assembly and also the content of certain nuclides in each sample. based on the published information, two depletion models describing depletion of istc fuel were made. one model was made in the triton sequence and one was made in the origen-arp sequence. the model for comparison was similar to the model described in previous section, only the geometry and composition were modified for description of the fuel assembly from the project [5, 6]. 2.3. influence of uncertainties 2.3.1. effect on the spent fuel inventory study of the influence of uncertainties was performed on changes in a content of group of nuclides in the spent fuel inventory. these nuclides were determined using the tsunami sensitivity calculation which determined that said nuclides affect reactivity considerably. the sensitivity calculation used spent fuel inventory determined by the model described in the section 2.1. the inventory was loaded into the tsunami model, which was obtained from the depletion model by modification of depletion parameters to sensitivity control parameters. determined nuclides are listed in table 2. nuclides listed in this table had the minimal sensitivity of 10−4 in at least one material. three nuclides marked with the star had smaller sensitivity, however they were accounted for because of their significant inaccuracy determined by the comparison with experimental data. effect of the uncertainties on the spent fuel inventory was studied by comparison of the inventory calculated without any applied uncertainty and the inventory calculated with one arbitrary uncertainty applied. uncertainties used in this comparison include changes of parameters describing the operation of the reactor, fuel assembly geometry or composition. 15 t. czakoj, j. frybort, m. lovecky acta polytechnica ctu proceedings quantity nominal lower higher moderator temperature 555 k 541 k 573 k fuel temperature 885 k 785 k 985 k boron content 524 ppm 0 ppm 1049 ppm power level 30.535 g cm−3 27.4815 g cm−3 33.5885 g cm−3 table 3. operational uncertainties uncertainty nominal lower higher absorber content (wt) 3.35 % 3.2 % 3.5 % uranium content 1230/1141 g 1213/1124 g 1247/1158 g enrichment (wt) 3.6/4/4.6 % 3.55/3.95/4.55 % 3.65/4.05/4.65 % outer fuel diameter 0.78/0.76 cm 0.777/0.757 cm inner fuel diameter 0/0.12 cm 0/0.09 cm outer cladding radius 0.91 cm 0.906 cm 0.914 cm inner cladding radius 0.793/0.773 cm 0.799/0.779 cm table 4. material uncertainties element content [%] element content [%] n 2,9e-03 pb 8,0e-04 f 2,2e-02 zn 1,1e-03 na 5,2e-03 cu 1,1e-03 mg 7,0e-03 ni 1,4e-02 si 7,1e-03 co 2,0e-05 cr 3,0e-03 fe 1,6e-02 mn 2,0e-04 table 5. content of impurities [7] complete list of the used uncertainties is presented in tables 3, 4 and 5. all the calculated inventories had equal burnup of 50.4 mwd kg−1hm, which corresponds to the burnup of spent fuel after five years depletion in the nuclear reactor. because of the preservation of the burnup, cases with lower or higher power level were depleted for longer or shorter time, respectively. 2.3.2. effect on the multiplication factor effect of the uncertainty on the multiplication factor was evaluated by determination of the multiplication factor of spent fuel. the composition of this fuel was calculated in previous part. the evaluation was made in keno-vi 3d stochastic sequence. the model was created by loading of the composition to the gd-2m+ depletion model and its modification for keno-vi sequence. the calculation used 2000 neutron generations with 50000 neutrons per generation and first 50 generations being inactive. the calculation was automatically terminated once the stochastic deviation decreased below 0.00020 because of the limited available computation time. 3. results and discussion 3.1. comparison with experimental data calculated content of selected nuclides in each sample was compared to the experimentally determined content. this comparison was created as a fraction of computed and experimental data, marked as c/e. an arithmetic mean of c/e for every nuclide is in table 6 and in figure 2. as presented in table 6, the results are relatively accurate for most of the elements, such as uranium, cesium or neodymium, but for some elements the inaccuracies are larger, such as for 108pd or 151eu. at the moment, there are no clear reasons for these inaccuracies and more detailed research for clarification of these inaccuracies is needed. 3.2. influence of uncertainties 3.2.1. effect on spent fuel inventory figure 3 depicts changes in the spent fuel inventory after the modification of moderator temperature. results show that the biggest changes in the content are for the isotopes with high cross section in thermal energies. this phenomenon is caused by the worsened moderation, which leads to the worsened absorption in thermal part of the spectrum. changes in the spent fuel inventory after modification of fuel temperature are presented in figure 4. the effect of change of fuel temperature is smaller in comparison with the effect of change of moderator temperature, but is important to take account that change in fuel temperature is usually many times higher than changes in moderator temperature. the change in inventory after the modification of fuel temperature is mainly caused by change of doppler effect. higher temperature decreases the maximum of the resonance, which results in higher resonance absorption. higher absorption at isotope of 238u then leads to higher production of actinides such as plutonium [8]. figure 5 shows results for changes of boron content in the moderator. it is obvious that higher concentration of boron leads to higher content of gadolinium in the spent fuel. this is caused by the worsened moderation which results in lower absorption in thermal part of the spectrum. 16 vol. 14/2018 calculation uncertainties in spent fuel inventory determination c/e c/e c/e nuclide triton origen-arp nuclide triton origen-arp nuclide triton origen-arp 234u 1.03 ± 0.09 1.03 ± 0.09 242cm 1.23 ± 0.14 1.24 ± 0.14 149sm 0.80 ± 0.67 0.80 ± 0.67 235u 0.96 ± 0.00 0.96 ± 0.00 244cm 1.23 ± 0.05 1.23 ± 0.05 150sm 1.07 ± 0.43 1.07 ± 0.43 236u 0.95 ± 0.00 0.95 ± 0.00 241am 1.17 ± 0.01 1.17 ± 0.01 151sm 1.09 ± 0.45 1.09 ± 0.45 238u 0.99 ± 0.00 0.99 ± 0.00 243am 0.91 ± 0.03 0.91 ± 0.03 152sm 1.05 ± 0.37 1.04 ± 0.37 238pu 1.19 ± 0.01 1.20 ± 0.01 133cs 0.99 ± 0.01 0.99 ± 0.01 155gd 1.13 ± 0.15 1.13 ± 0.15 239pu 1.23 ± 0.00 1.23 ± 0.00 134cs 0.94 ± 0.01 0.94 ± 0.01 95mo 0.49 ± 0.15 0.49 ± 0.15 240pu 1.16 ± 0.00 1.16 ± 0.00 135cs 1.15 ± 0.01 1.10 ± 0.01 105pd 1.13 ± 0.62 1.13 ± 0.62 241pu 1.20 ± 0.00 1.20 ± 0.00 137cs 0.99 ± 0.00 0.99 ± 0.00 108pd 2.67 ± 3.73 2.67 ± 3.73 242pu 1.14 ± 0.01 1.14 ± 0.01 151eu 0.39 ± 0.09 0.39 ± 0.09 109ag 0.70 ± 0.11 0.70 ± 0.11 143nd 1.05 ± 0.00 1.05 ± 0.00 153eu 1.12 ± 0.73 1.12 ± 0.73 101ru 1.06 ± 0.40 1.06 ± 0.41 145nd 1.02 ± 0.00 1.02 ± 0.00 154eu 1.22 ± 0.94 1.22 ± 0.94 237np 0.79 ± 0.01 0.79 ± 0.01 146nd 1.01 ± 0.00 1.01 ± 0.00 155eu 1.09 ± 0.68 1.09 ± 0.67 144ce 0.99 ± 0.15 0.99 ± 0.15 148nd 1.01 ± 0.00 1.01 ± 0.00 147sm 1.03 ± 0.36 1.03 ± 0.36 table 6. comparison of calculated and experimental data figure 2. comparison of calculated and experimental data figure 3. effect of change in moderator temperature on composition. similar observation can be seen in the content of samarium. both its isotopes 149sm and 151sm have a high microscopic cross section for absorption in thermal energies. similarly to gadolinium, higher boron figure 4. effect of change in fuel temperature on composition. content worsens moderation which causes lower absorption in samarium isotopes [2]. results for different power levels are shown in figure 6. the power level change affects only few isotopes, content of most of the isotopes remains almost un17 t. czakoj, j. frybort, m. lovecky acta polytechnica ctu proceedings figure 5. effect of change in boron content on composition. figure 6. effect of change in power level on composition. changed. these changes are mainly caused by longer depletion time, which allows for a higher accumulation of certain isotopes in the spent fuel, caused by the decay of their parent nuclides. another reason for the changes in content is connected to the higher production of short-lived fission products and its daughter isotopes for higher power levels, such as 135xe or 105rh [9]. figure 7 shows results of impact of the variation in the gd2o3 content in spent fuel. this variation affects mainly the gadolinium content in spent fuel inventory, yet it also leads to increased content of 161dy and 159tb isotopes, due to them originating from gadolinium according to equations (1) (3) [10]. effect on content of any other nuclides is negligible. 158gd + n → 159gd β − −−−−−→ 18,479h 159tb (1) 159tb + n → 160tb   β−−−−→ 72,3d 160dy + n → 161dy +n → 161tb β − −−−−→ 6,906d 161dy (2) 160gd + n → 161gd β − −−−−−→ 3,66min 161tb β − −−−−→ 6,906d 161dy (3) figure 8 shows results of impact of change in the uranium content. as expected, the higher initial content of uranium leads to higher final content of all figure 7. effect of change in gd2o3 content on composition. figure 8. effect of change in uranium content on composition. investigated nuclides, however, the difference between content of various nuclides is visible. the results of variations in the enrichment are depicted in figure 9. the changes in uranium isotopes content were expected because their initial content was also modified together with the enrichment [11]. variations in content of isotopes 242am, 244cm or 242pu are also observed. their content in spent fuel is higher with lower enrichment. this phenomenon is caused by their production from the 238u because the initial content of this isotope also varies with enrichment. another effect of change in enrichment is also connected to modification of neutron spectrum because lower content of 238u leads to lower parasitic absorption. figures 10 and 11 show results for changes in fuel and cladding radii. the impact of fuel radius change is negligible, which is probably caused by modification of fuel density made with the objective of preservation of fuel content. bigger effects observed for fuel pins without gadolinium are obviously caused by their higher number in the fuel assembly. effect of modification of outer cladding radius is higher, which is caused by change of the amount of moderator near to the fuel pin. this change leads to better (or worse) moderation and subsequently to a modification of the neutron spectrum. effect of variation of inner cladding radius is negligible because it only changes the amount of cladding material, which has only small influence on neutrons. 18 vol. 14/2018 calculation uncertainties in spent fuel inventory determination figure 9. effect of change in enrichment on composition. figure 10. effect of change in fuel radius figure 12 shows the effect of impurities in fuel. this effect is both positive and negative and for most applications, it can be considered negligible. nevertheless it is possible to find some cases, in which the impurities can be important. for example, content of nitrogen can lead to production of radioactive 14c, which may have impact on the overall activity [12]. 3.2.2. effect on multiplication factor effect of the operational uncertainties was tested only with investigated uncertainty included in the depletion simulation. this means that the determination of multiplication factor is not influenced by presence of the uncertainty, but it is influenced only by change in inventory. results of this calculation can by seen in figure 13. the biggest impact definitely comes from the change of boron content in the moderator. much smaller, but also important is the effect of change in the moderator temperature. other uncertainties are of smaller importance. both this effects are mainly caused by modification of content of nuclides with high cross section for absorption such as 155gd, 157gd, 149sm and 151sm which negatively influence neutron balance [2]. figure 14 shows the influence of the material uncertainties on the multiplication factor. this influence has smaller importance than the operational uncertainties. the most important uncertainty is initial enrichment of the fuel, which will be mainly caused by change in final content of fissile isotope 235u. second figure 11. effect of change in cladding radius figure 12. effect of impurities in the fuel biggest effects comes from different initial uranium content, which can probably mainly assigned to change in boron content together with changed content of fissile isotopes of plutonium and uranium. figure 15 shows impact of changes in geometry. these uncertainties have only small importance and the most significant is the change in the outer diameter of fuel with gadolinium absorber together with the change in its inner cladding diameter. 4. conclusions two depletion models of gd-2m+ fuel were created and the triton model was optimized for faster calculation. this model was used for the study of influence of the operational and manufacturing uncertainties. comparison of computed and experimental data was conducted for both used depletion sequences. no important difference in results from both sequence was found. obtained results were relatively accurate, but few nuclides, such as 108pd or 151eu, with inaccurate determination were found. the dependence between the inaccuracy of 108pd or 151eu and between the group of tested uncertainties was not found. on the contrary, the sensitivity of 242cm and 244cm content to the boron content in the moderator and to the moderator temperature was found, which is caused by modification of neutron spectrum. the influence of uncertainties on the spent fuel inventory and the multiplication factor was investigated. the determination of the multiplication factors has 19 t. czakoj, j. frybort, m. lovecky acta polytechnica ctu proceedings figure 13. effect of operational uncertainties on multiplication factor only with uncertainty included in fuel depletion figure 14. effect of material uncertainties on multiplication factor allowed for a comparison of importance of each uncertainty. it was found that most important are the operational uncertainties, especially boron content in the moderator and the moderator temperature. although fuel temperature change seems to be much low important than the previously mentioned, it is important to remember that fuel temperature changes are much high than the moderator temperature changes. importance of other investigated operational uncertainties is much lower than the importance of the previous. furthermore, it was discovered that in order to apply the conservative approach, it is necessary to set up lower boron content and higher temperature of moderator. this approach leads in faster depletion of gadolinium burnable absorber which leads in lower absorption. on the contrary, higher fuel temperature leads in higher gadolinium content, but this effect is prevailed by higher production of fissile isotopes of plutonium together with lower 235u depletion. due to this fact, higher fuel temperature is more conservative. effect of all manufacturing uncertainties is lower than the most important operational uncertainties. the most important is correct set up of the enrichment and the uranium content in fuel. it is conservative to overestimate both values. references [1] b. t. rearden, m. a. jessee. scale code system 6.2.1. tech. rep., 2016. doi:10.2172/1326509. figure 15. effect of uncertainties of geometry on multiplication factor [2] m. chadwick, m. herman, p. obložinský, et al. endf/b-vii.1 nuclear data for science and technology: cross sections, covariances, fission product yields and decay data. nuclear data sheets 112(12):2887 – 2996, 2011. special issue on endf/b-vii.1 library, doi:10.1016/j.nds.2011.11.002. [3] p. mikolas, z. fakezasova, m. sasek, et al. zmena casti ppbz pro zavedeni paliva gd-2m+edu: 4.3 jaderne charakteristiky, 2013. [4] j. leppanen, m. pusa, e. fridman. overview of methodology for spatial homogenization in the serpent 2 monte carlo code. annals of nuclear energy 96(supplement c):126 – 136, 2016. doi:https://doi.org/10.1016/j.anucene.2016.06.007. [5] s. pavlov. investigation of the vver-440 fuel burnup and nuclide composition for the use in burnup credit-related tasks (buc-ii) final scientific technical report on task 5, dimitrovgrad. istc project #3958p, dimitrovgrad, 2011. [6] l. markova, f. havluj. specification of simplified benchmark based on #3958 istc pie. nuclear research institute at rez, czech republic, 2011. [7] l. pajo. uo2 fuel pellet impurities, pellet surface roughness and n (18o)/n(16o) ratios, applied to nuclear forensic science. ph.d. thesis, university of helsinki, faculty of science, department of chemistry, 2011. [8] j. r. lamarsh. introduction to nuclear engineering, 2nd edition. addison-wesley, 1983. [9] b. hermansky. dynamika jadernych reaktoru : docasna vysokoskolska ucebnice pro studenty obor jaderne inzenyrstvi. ceskoslovenska redakce vn mon, praha, 1st edn., 1987. vydavatel: csr. ministerstvo skolstvi, mladeze a telovychovy. [10] y. tanoue, t. yokoyama, m. ozawa. resource evaluation of heavy rare earth derived from the spent gd2o3 burnable poison in lwrs. journal of energy and power engineering 10(4), 2016. doi:10.17265/1934-8975/2016.04.004. [11] s. cerne, o. hermann, r. westfall. reactivity and isotopic composition of spent pwr (pressurized-waterreactor) fuel as a function of initial enrichment, burnup, and cooling time. tech. rep., 1987. doi:10.2172/5561567. [12] m.-s. yim, f. caron. life cycle and management of carbon-14 from nuclear power generation. progress in nuclear energy 48(1):2 – 36, 2006. doi:https://doi.org/10.1016/j.pnucene.2005.04.002. 20 http://dx.doi.org/10.2172/1326509 http://dx.doi.org/10.1016/j.nds.2011.11.002 http://dx.doi.org/https://doi.org/10.1016/j.anucene.2016.06.007 http://dx.doi.org/10.17265/1934-8975/2016.04.004 http://dx.doi.org/10.2172/5561567 http://dx.doi.org/https://doi.org/10.1016/j.pnucene.2005.04.002 acta polytechnica ctu proceedings 14:8–14, 2018 1 introduction 2 calculation methods 2.1 depletion model 2.2 comparison with experimental data 2.3 influence of uncertainties 2.3.1 effect on the spent fuel inventory 2.3.2 effect on the multiplication factor 3 results and discussion 3.1 comparison with experimental data 3.2 influence of uncertainties 3.2.1 effect on spent fuel inventory 3.2.2 effect on multiplication factor 4 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0032 acta polytechnica ctu proceedings 12:32–37, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app microsleeps and their detection from biological signals martin holub∗, martina šrutová, lenka lhotská czech technical university in prague, faculty of electrical engineering, department of cybernetics, karlovo nám. 13, praha 2, 121 35, czech republic ∗ corresponding author: martinholub@mail.com abstract. microsleeps (ms) are a frequently discussed topic due to their fatal consequences. their detection is necessary for the purpose of sleep laboratories, where they provide an option for the quantifying rate of sleep deprivation level and objective evaluation of subjective sleepiness. many studies are dealing with this topic for automotive usage to design a fatigue countermeasure device. we made a research of recent attitude to the development of the automated ms detection methods. we created an overview of several ms detection approaches based on the measurement of biological signals. we also summarized the changes in eeg, eog and ecg signals, which have been published over the last few years. the reproducible changes in the entire eeg spectrum, primarily with the increased activity of delta and theta, were noticed during a transition to fatigue. there were observed changes of blinking rate and reduction of eye movements during the fatigue tasks. ms correspond with variations in the autonomic regulation of the cardiovascular function, which can be quantified by hrv parameters. the decrease in hr, vlf, and lf/hf before falling asleep was revealed. eeg signal, especially its slow wave activity, considered to be the most predictive and reliable for the level of alertness. in spite of the detection from eeg signal is the most common method, eog based approaches can also be very efficient and more driver-friendly. besides, the signal processing in the time domain can improve the detection accuracy of the short events like ms. keywords: microsleep, automatic detection, detection methods, eog, eeg, ekg. 1. introduction microsleeps (ms) are mostly detected for the purpose of automotive usage and sleep laboratories. the evaluation of ms stages in the sleep laboratories provides an option for the quantifying rate of sleep deprivation level and an objective assessment of subjective sleepiness. they use the definition of microsleep (ms), which arose as a consequence of the evaluation method of the psg (polysomnography) or mslt (multiple sleep latency test) data, where each 30-second epoch is evaluated separately. to score this epoch as a sleep, more than 15 seconds should show some sleep activity [1]. episodes with sleep symptoms, but with a short duration between 3 and 15 seconds are defined as the microsleeps. the epoch would be classified as a sleep in case, they were longer than 15 seconds [2]. the theta rhythm with slow eye movements, k complexes, or sleep spindles is included. microsleeps can be accompanied by the same behavioural changes as during the states of drowsiness. in spite of the short duration of microsleep, it can be inconvenient and dangerous for many people and for their surroundings to perform monotonous activity while being tired and drowsy. many studies are dealing with this topic for automotive usage to design a fatigue countermeasure device. for those kinds of studies, it isn’t important to distinguish the events of drowsiness, microsleeps or fatigue correctly by the terminology, because all the states associated with the loss of attention are dangerous in the traffic. in the studies observed, the mentioned lapses of responsiveness are often manifested by absent or prolonged responses and short suspension of performance. these symptoms are accompanied by droopy eyes, slow eyelid closure, head nodding and many changes in biological signals. we aimed to make a research of recent attitudes for the development of the automated ms detection methods. we created an overview of several ms detection approaches based on biological signals measurement. they are mostly based on the processing eeg (electroencephalography) signals or eeg accompanied by eog (electrooculography) signals. we also summarized the changes in the basic biological signals like eeg, eog and ecg (electrocardiography), which have been published over the last few years. 2. signal changes 2.1. changes in eeg signals the changes in the entire eeg frequency spectrum were noticed during transition to fatigue [3–8]. the increased activity of delta and theta rhythms was primarily observed. it could reflect decreased cortical arousal during a monotonous task. the removing noise from the eeg is crucial for observation of right changes. the filtration method is especially necessary 32 http://dx.doi.org/10.14311/app.2017.12.0032 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 microsleeps and their detection from biological signals for the low frequency signal, which is influenced by artefacts from activities such as breathing and movement. slow activity is prone to be also contaminated by eye movement artefact. various areas of the brain behaved differently during fatigue, during its onset delta and theta activity were present mostly in the frontal, central, and parietal areas of the brain with some anterior alpha and posterior beta [3, 5]. the occipitoparietal alpha spread to anterior areas and become more centrofrontal and temporal alpha during drowsiness [3]. the reproducibility of the eeg magnitude changes in the delta, theta, alpha and beta bands during drive fatigue was tested [9]. there was revealed high reproducibility for the delta and theta bands [9]. during the transitional phase to fatigue, there were no significant differences in the delta and alpha magnitudes across the entire brain [9]. the inter-hemispheric eeg coherence analysis was shown in the study of jap et al. [10], which revealed that the interhemispheric coherence level was significantly higher at the frontal and occipital sites compared to the central, parietal, and temporal sites throughout the driving sessions. 2.2. changes in eog signals due to the rich connectivity between eyes and brain, different eog analyses can be used to identify the states similar to fatigue. the changes of the blinking rate were observed during the fatigue tasks in the published papers [5]. the work of lal [5] showed, that the fast eye movements and conventional blinks during alert were replaced by no eye movements and small fast rhythmic blinks during fatigue, and that the slow blinks were present during deeper drowsiness for all the subjects. wierwille et al. [11] declared that slow eyelid closure is a reliable estimate of the level of drowsiness. in the study of papadelis et al. [12], a significant increase of eye blinks’ number and duration before driving errors was observed. 2.3. changes in ecg signals the stages of microsleeps, drowsiness or fatigue correspond with the changes in the autonomic regulation of the cardiovascular function. the hrv (heart rate variability) parameters are used to quantify these changes. the gradual and sustained decrease in the vlf range minutes before falling asleep was observed [13]. the ratio lf/hf, which corresponds with the sympathovagal balance, was very low compared to baseline wake values for about 5 minutes before the event [13]. it is a sign of the relaxation taking over. there is a significant surge in this balance after the falling asleep due to the increased stress [13]. heart rate decreased significantly during falling asleep event [13] and fatigue (by 7 bpm) [5]. same changes in the heart rate were observed during prolonged and monotonous driving [14]. also, the rr variability declined over the falling asleep events in the study of furman [13]. 3. detection approaches in the following section, there are presented and compared (table 1) detection methods, which have been issued during the recent years and which were dealing with the states similar to microsleeps. 3.1. detection by vuckovic et al. (published 2002) the classification of the drowsy states from full spectrum eeg recordings was published in the paper [15]. parameterization was based on a moving cross-spectral density csd. the interhemispheric csd was calculated from homologous contralateral electrodes [15]. the intrahemispheric csd was calculated from electrodes on the left hemisphere [15]. the csd was computed for the signal lasting one epoch and the signal from the same electrode delayed by 50% of the epoch [15]. the main reason to choose interhemispheric csd was that the correlation of homologue electrodes decreases in drowsiness [16]. as well as the intrahemispheric csd detects the transition from alertness to drowsiness and vice versa because of the changes in the synchronization of the intrahemispheric eeg signal [16]. a convolutional neural network, auto encoder, and linear dynamic system described accurately in [15] were used for classification. 3.2. detection by golz et al. (published 2007) different ms detection approaches and their combinations were presented in the paper [17]. the adaptive signal processing was based on the signals of the brain electric activity, variation in the pupil size, and eyelid movements. totally fifteen signals had been acquired, seven signals of eeg, vertical and horizontal eog channel, and six signals from eye tracking system (the pupil size and two coordinates of eye gaze on the plane of projection). the features were firstly extracted in the frequency domain by one linear method of modified periodogram and then in the state space by using the nonlinear method of the delay vector variance (dvv) [18]. the results show, that periodogram was more powerful for the feature extraction than the dvv method and that the complementing power spectral density by dvv features showed only small improvements. the achieved features were used as input vectors in automatically learning classification algorithm, where the decision process was based on the support vector machines and three learning vector quantization neural networks. all algorithms were optimized to gain maximal classification accuracy. the best results were achieved by the fusion of features for all available signals. proposed methods were created with respect to the future online driver monitoring. [17] 33 m. holub, m. šrutová, l. lhotská acta polytechnica ctu proceedings figure 1. the table shows the comparison of the published methods. 3.3. detection by furman et al. (published 2008) this study [13] checked a feasibility of the new detection approach based on the processing of the ecg signal, where the trends of heart rate variability (hrv) were observed. the hrv parameters were used to look into the changes in the autonomic regulation of the cardiovascular function. physiological changes were quantified before, during and after an event of microsleep. it was found that there was a gradual and sustained decrease in the vlf range minutes before falling asleep events. the sympathovagal balance, which was represented by lf/hf, was very low compared to baseline wake values for about 5 minutes before the events. it was a sign of relaxation taking over. after falling asleep there was a surge in this balance signifying an increased stress aimed to overcome the drowsiness. the mean hr and rr variability decreased during falling asleep events, thus confirm the autonomic changes. during a microsleep, there was a decrease in sympathetic drive with increased vagal activity. [13] 3.4. detection by tsai et al. (published 2009) real-time system [19] for detecting a drowsy driver was based on eeg signal processing. six channels of eeg were acquired by a novel active dry electrode non-invasive system. the wavelet transform using daubechies 2 wavelet was applied on the each channel. integrated eeg and zero crossing features were computed from the scales corresponding to the frequency of beta, alpha and theta rhythms. all 36 parameters were used as an input vector to the artificial neural network with 18 hidden nodes and two output nodes. system rejected the detection accuracy of drowsy subject 90.9% with too many warnings signals while the subject was totally alert. [19] 3.5. detection by garcés et al. (published 2010) the automatic detector of drowsiness episodes [20] was based on the processing of one eeg channel. the butterworth band-pass filter with cut-off frequencies of 0.5 and 60 hz was used for signal pre-processing 34 vol. 12/2017 microsleeps and their detection from biological signals and then segmented into the blocks of duration of 1, 2, 5 and 10 seconds. the length of 5 seconds was considered the most appropriate to take the decision in the real time. twelve features were computed from spectral analysis and wavelet decomposition of the eeg signal: the central frequency, the first quartile frequency, the maximum frequency, the total energy of spectrum, the power of theta and alpha bands, the number of zero crossing and the integrated eeg for scales 3, 4 and 5 of daubechies 2 wavelet transform. in this case, the equivalent filters of mentioned scales had approximately same frequency as beta, alpha and theta rhythms. twenty-five neural networks with three layers (first 12 neurons, last one neuron with tan-sigmoid function, hidden-between 5 and 30 neurons) were tested for classification. the best architecture was neural network (12-20-1), for which 81.7% of drowsiness was detected. [21] 3.6. detection by peiris et al. (published 2011) the detection of lapses in responsiveness [22] was based on the signal processing of the electroencephalogram. the features using spectral analysis (mean spectral power, normalized spectral power and spectral power ratios) and three nonlinear methods (fractal dimension, approximate entropy and lempel-ziv complexity) were calculated and compare to determine their efficacy in eeg-based detection of lapses. principal component analysis was used to reduce the dimensionality of the feature matrix. the detector model created using mean spectral power features achieved the best failure state estimation performance. there were reached sensitivity 73.5% and selectivity 25.5%. among the nonlinear approaches the detector based on the lempel-ziv complexity showed the highest performance. the performance of the detector was no greater by complementing spectral power features by lempel-ziv complexity. it was similar to the finding of [17]. the use of meta-learner weights showed a marginally higher performance than the use of constrained least-squares metalearner weights. [22] 3.7. detection by zhu et al. (published 2014) the drowsiness detection [23] was based on two electrooculography signals, vertical and horizontal. these signals were filtered, saturated and normalized during pre-processing. the automatic detection technic of slow eye movements [21, 24, 25] based on wavelet transform was used for the feature detection. the features were slow eye movements number, duration of closing phase of blink, peak and mean velocities in closing and opening phases of blink. the wavelet transformation was also used for the extraction of two other features ratio of low and high frequency on both horizontal and vertical eogs. the linear dynamic system was used for post processing. the convolution neural network, which was described in detail in the paper [23], was used for classification. 3.8. detection by holub et al. (published 2015) ms classifier [26] was based on the independent analysis of each couple of eog signals in the time domain. the basic idea of the algorithm was to observe differences between the minimum and maximum values in the appropriately chosen sliding window, which was gradually applied to the entire length of both eog signals. by merging information from the signals of the left and the right eye was calculated the final feature, which was able to appropriately distinguish ms events from the rest of the awake signal [26]. this parameter highlighted ms states, which were manifested in the signal as the episodes without significant eye movements and fast changes. these episodes were terminated by the specific high-frequency wave, which corresponded to blink and arousal from the ms. the final parameter was thresholded to distinguish ms from the rest of the signal. the differences between fix and adaptable thresholding were presented. the fix threshold was appropriately set as the same value for all signals, while the calculation of adaptable threshold was based on the signal similarity of ms episodes with sleep stages nrem1 and nrem2, which were used to set this limit value. there were reached precision 67% and recall 82% in the method of adaptable thresholding. 4. conclusions eeg signal has been considered by the researchers as the most predictive and reliable for the level of alertness. previous studies confirmed that detection of changes in slow wave activity might form the basis of a fatigue countermeasure device. these approaches use mostly wavelet transform or methods of periodogram, which can be suitable for the detection of drowsy states, but they can provide the insufficient time resolution for detection of very short events like ms. to achieve a good frequency resolution, it is necessary to analyse the longest possible part of the signal. but the time resolution deteriorates with increasing length of the sliding window, which is applied to the analysed signal. this fact is undesirable due to the ms detection of an event that may take only 3 seconds. the window has to be always chosen long enough with respect to the stationarity of the signal and frequency resolution. the same problem of insufficient time resolution can be in the detection from ecg, where the heart rate variability is counted. in comparison to detection in the frequency domain, the signal processing in the time domain can provide higher accuracy and faster response of the detector. in spite of the detection from eeg signal, it is the most common method, eog based approaches can also be very efficient and more driver-friendly. the electrodes for the measuring of eog signals can be 35 m. holub, m. šrutová, l. lhotská acta polytechnica ctu proceedings easily implemented for example to headbands, hats or glasses, which provide unlike measuring multi-channel eeg greater comfort and easy application for the drivers. acknowledgements this work has been supported by the project grant application no. sgs16/231/ohk3/3t/13 of the czech technical university in prague. references [1] c. iber. the aasm manual for the scoring of sleep and associated events: rules, terminology and technical specifications. american academy of sleep medicine, 2007. 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[20] a. g. correa, e. l. leber. an automatic detector of drowsiness based on spectral analysis and wavelet decomposition of eeg records. in 2010 annual international conference of the ieee engineering in medicine and biology, pp. 1405–1408. ieee, 2010. [21] e. magosso, m. ursino, a. zaniboni, et al. visual and computer-based detection of slow eye movements in overnight and 24-h eog recordings. clinical neurophysiology 118(5):1122–1133, 2007. doi:10.1016/j.clinph.2007.01.014. [22] m. t. peiris, p. r. davidson, p. j. bones, r. d. jones. detection of lapses in responsiveness from the eeg. journal of neural engineering 8(1):016003, 2011. [23] x. zhu, w.-l. zheng, b.-l. lu, et al. eog-based drowsiness detection using convolutional neural networks. in ijcnn, pp. 128–134. 2014. 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[25] e. magosso, m. ursino, a. zaniboni, e. gardella. a wavelet-based energetic approach for the analysis of biomedical signals: application to the electroencephalogram and electro-oculogram. applied mathematics and computation 207(1):42–62, 2009. doi:10.1016/j.amc.2007.10.069. 36 http://dx.doi.org/10.1016/j.sleep.2006.06.011 http://dx.doi.org/10.1097/00004691-198710000-00002 http://dx.doi.org/10.1017/s0048577201393095 http://dx.doi.org/10.1016/0013-4694(87)90096-4 http://dx.doi.org/10.1016/j.ijpsycho.2004.07.001 http://dx.doi.org/10.1016/0001-4575(94)90019-1 http://dx.doi.org/10.1016/j.clinph.2007.04.031 http://dx.doi.org/10.1016/s1350-4533(02)00030-9 http://dx.doi.org/10.1177/155005949602700207 http://dx.doi.org/10.1007/s11265-007-0083-4 http://dx.doi.org/10.1016/j.clinph.2007.01.014 http://dx.doi.org/10.1016/j.medengphy.2006.01.002 http://dx.doi.org/10.1016/j.amc.2007.10.069 vol. 12/2017 microsleeps and their detection from biological signals [26] m. holub, m. šrutová, l. lhotská. adaptable microsleep detection based on eog signals: a feasibility study. in computational intelligence for multimedia understanding (iwcim), 2015 international workshop on, pp. 1–4. ieee, 2015. 37 acta polytechnica ctu proceedings 12:32–37, 2017 1 introduction 2 signal changes 2.1 changes in eeg signals 2.2 changes in eog signals 2.3 changes in ecg signals 3 detection approaches 3.1 detection by vuckovic et al. (published 2002) 3.2 detection by golz et al. (published 2007) 3.3 detection by furman et al. (published 2008) 3.4 detection by tsai et al. (published 2009) 3.5 detection by garcés et al. (published 2010) 3.6 detection by peiris et al. (published 2011) 3.7 detection by zhu et al. (published 2014) 3.8 detection by holub et al. (published 2015) 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0072 acta polytechnica ctu proceedings 22:72–76, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app electrical properties of fly ash geopolymer composites with graphite conductive admixtures cecílie mizerová∗, ivo kusák, pavel rovnaník brno university of technology, faculty of civil engineering, žižkova 17, 602 00 brno, czech republic ∗ corresponding author: mizerova.c@fce.vutbr.cz abstract. construction materials with increased electrical conductivity could be possibly used in health monitoring of structures (stress, deformation, damages), their maintenance or traffic monitoring. the aim of this study was the application of functional filler and its influence on the electrical properties of the alkali-activated fly ash matrix. the graphite powder was added to the reference material in the amount of 2–10 %. besides the assessment of the critical amount of filler necessary to achieve a percolation threshold in the structure of the composite, the effect on the electrical properties of the matrix (resistance, capacitance, conductivity) was determined. the optimal amount of filler was also determined with respect to the changes in microstructure of the binder and its mechanical properties. keywords: fly ash, geopolymer, graphite, electrical properties, mechanical properties, microstructure. 1. introduction recently, one of the field in materials and structures research is the concept of advanced smart (or selfsensing) material. its composition design is based on the content of conductive fillers that enhance its electrical properties. for example, smart composite with higher electrical conductivity can be used for the assessment of structural health in terms of stress monitoring, localization of cracks and damages, humidity and temperature changes etc. [1]. there is a wide range of different functional fillers to use: steel fibres, carbon black, graphite powder, carbon fibres (cf), carbon nanotubes (cnf), graphene, nickel powder. so far, this approach has been studied predominantly in traditional cement concretes and composites technology. alkali-activated materials are binders derived from the reactions of raw materials with a high content of amorphous aluminosilicate phase with alkaline activator solution. alkaline activator (most often in form of alkali hydroxides and/or alkali silicates) is required for the appropriate conditions for the dissolution rate and the formation of new phases [2]. the first theoretical background of alkaline activation was developed by glukhovsky in 1959 [3]. the term “geopolymer” was firstly used by prof. davidovits and refers to a material formed by almost exclusively highly coordinated units of aluminosilicate structures forming regular polymeric three-dimensional structure [4]. compared to clay-based geopolymers, fly ash provides more favourable rheology of the fresh binder [5], strength of the hardened binder can be enhanced by thermal treatment. properties of alkali-activated binders are comparable or superior to portland cement binders [6]. fly ash geopolymers are generally more durable because of higher chemical and thermal resistance [7]. moreover, the geopolymer production using mostly secondary raw materials do not require an exploitation of natural resources and high energy consumption that significantly contribute to lower emissions of co2 compared to portland cement [8]. according to hanjitsuvan et al. [9, 10], the electrical conductivity of fly ash geopolymer matrix is affected by the naoh concentration in activator solution, the frequency spectrum and liquid activator/ash ratio (l/a). payakaniti [11] studied properties of geopolymer incorporated with cf stating the percolation threshold was reached by the geopolymer with 0.5% cf content, this value was increasing with growing l/a. this cf content was also responsible for the best mechanical performance. vaidya et al. [12] tested fly ash geopolymer concrete beams and cylinders with 0.4% cf while exposed to bending and compressive stress to assess the changes of electrical resistance. regarding the application of cnt in geopolymer concrete, saafi et al. [13] states that the optimal dosage is 0.5% which ensures the improvement of both electrical and mechanical performance, higher percentage leads to less effective dispersion and thus cluster formation within the matrix. in another study, the author used graphene as a conductive filler [14]. only 0.35% addition of graphene resulted in significant increase of strength and modulus of elasticity accompanied by lower porosity and 209% increase of conductivity. this paper is focused on the application of graphite powder as a conductive admixture in fly ash geopolymer mortars considering its impact on selected electrical properties, strength and changes in binder microstructure. 2. materials and methods the geopolymer binder was produced by mixing of fly ash (fa) with commercial sodium silicate solution with sio2/na2o = 1.6. the fly ash is a product of black coal combustion and its chemical composition is given in table 1. quartz sand with a maximum 72 http://dx.doi.org/10.14311/app.2019.22.0072 http://ojs.cvut.cz/ojs/index.php/app vol. 22/2019 electrical properties of fly ash geopolymer composites. . . sio2 al2o3 fe2o3 cao mgo stotal na2o k2o (%) 49.82 24.67 7.5 3.91 2.68 0.91 0.7 2.78 table 1. chemical composition of fly ash. ref g2 g4 g5 g6 g8 g10 fly ash (g) 350 350 350 350 350 350 350 sodium silicate (g) 280 280 280 280 280 280 280 graphite powder (g) 7 14 17.5 21 28 35 sand (g) 1050 1050 1050 1050 1050 1050 1050 2% triton x-100 (g) 7 14 17.5 21 28 35 1% lukosan (g) 3 6 7.5 7.5 7.5 7.5 water (g) 35 30 30 32 32 32 32 table 2. mix composition of fa geopolymer mortars with graphite powder. grain size of 2.5 mm was used as aggregate. graphite powder pmm 11 was used as a conductive filler in 2, 4, 5, 6, 8 and 10% wt. of the fly ash and its dispersion within the matrix was supported by the addition of dispersing agent (2% triton x-100). defoaming agent (1% lukosan) was introduced to reduce the air content in fresh and hardened binder. the raw mix composition is presented in table 2. each of the mixes was produced following these steps: at first, the fly ash was mixed with water glass. then, the graphite powder and dispersing agent were added and homogenized with the binder and a small amount of water, if needed. thereafter, each fraction of aggregate (from fine to coarse) was added and mixed properly. the defoaming agent and remaining amout of water were added at the very end of the mixing procedure. the mortars were cast into prismatic moulds (40 × 40 × 160 mm) and covered with a plastic sealant to avoid moisture loss. after 2 hours in the ambient conditions, the specimens were heated at 40 °c for 24 h. after demoulding, the hardened specimens were stored in the laboratory conditions (22 ± 2 °c, ϕ = 45 ± 5%) till the age of testing. the electrical and mechanical properties (compressive and flexural strength) were tested at the age of 56 days. the prepared prismatic samples were characterized by impedance spectroscopy in the range of 40 hz to 1 mhz using an agilent 33220a sinusoidal signal generator and an agilent 54645a dual-channel oscilloscope. the output voltage of the signal generator was 5.5 v. the input values for the electrical capacity and the resistance of the oscilloscope were 13 pf and 1 mω, respectively. in order to perform impedance analysis, the prismatic specimens were placed between parallel brass electrodes (30 × 100 mm) so that a distance between electrodes was 40 mm. conductivity measurements were performed in the range of 10 to 3510 mhz. the microstructure and porosity of the geopolymers were evaluated using micromeritics poresizer 9310 and scanning electron microscope (sem) tescan mira3 xmu. 3. results and discussion electrical resistance of all samples decreased with higher frequency applied, as can be seen from fig. 1. the lowest resistance was observed in case of 10% graphite content (g10), at all frequencies. the differences in resistance curves are clearly visible up to 10 khz. above this frequency, the resistance of all samples fluctuate and the curves overlap. at 100 hz, the resistance of reference sample was 159.80 mω. the resistance of g5 dropped to 43.78 mω and 0.85 mω only at g10. changes of resistance and conductivity at chosen frequency are shown in fig. 2. while the resistance decreased continuously, we registered an exceptional increase in conductivity of specimen g10 from 0.38 to 6.31 ms·m−1. these results suggest that geopolymer with 10% of graphite may have reached the percolation threshold. the capacitance of geopolymers with various graphite content as a function of frequency is shown in fig. 3. up to 6% of graphite, the changes in capacitance are comparable, but the capacitance of g8 and g10 shows the most notable increase over all frequencies. capacitance of all samples except g10 decreased with higher frequency. the capacitance of g10 first dropped from 222 to 122 pf (40–180 hz), within the frequency 180–10 000 hz it was about 125 pf and at higher frequency applied it was decreasing again. the changes of mechanical performance depending on the graphite content are displayed in fig. 4. the flexural strength was not negatively influenced by the graphite addition and fluctuated around 5 mpa at all graphite concentrations. on the contrary, the compressive strength gradually decreased. while the reference sample without graphite reached 23.5 mpa, the samples with 2% graphite content (g2) reached 17.7 mpa and 10% graphite content (g10) 73 c. mizerová, i. kusák, p. rovnaník acta polytechnica ctu proceedings figure 1. variation in electrical resistance of fa geopolymers with different graphite content. figure 2. resistance (at 100 hz) and conductivity (at 100 mhz) of fa geopolymers with different graphite content. figure 3. variation in capacitance of fa geopolymers with different graphite content. 74 vol. 22/2019 electrical properties of fly ash geopolymer composites. . . figure 4. mechanical properties of fa geopolymers with different graphite content. figure 5. cumulative intruded volume of fa geopolymers with different graphite content. (a) . ref. (b) . g2. (c) . g5. (d) . g10. figure 6. sem images of ref, g2, g5 and g10 geopolymers. 75 c. mizerová, i. kusák, p. rovnaník acta polytechnica ctu proceedings only 7.8 mpa, which is equal to 25% and 65% drop of strength, respectively. regarding the microstructure of the geopolymer binder, even 2% of graphite caused an increase in total porosity of the samples (fig. 5). while the geopolymers with a graphite content of 2–8% did not exhibit considerable differences, we observed a further increase in pore volume of g10 which is more than 60% higher than the reference sample. in fig. 6 we can see the nature of chosen binders in sem images. reference sample is characterized by homogeneous amorphous binder phase whereas number of pores and inhomogeneities increased with the amount of conductive filler. 4. conclusions the results mentioned above confirm that graphite powder can be applied to fly ash geopolymer matrix to enhance its electrical properties. higher graphite content resulted in reduced electrical resistance, especially in frequency up to 1 000 hz. likewise, the samples with high graphite content exhibited an increase in capacitance. the best performance in terms of electrical properties was observed in case of geopolymer with 10% graphite content providing significant increase in conductivity, which may refer to achieving the percolation threshold. however, the presence of conductive filler in such concentration deteriorated the mechanical performance in compression by more than 65%; flexural strength of all samples remained constant regardless of the amount of graphite powder. despite the use of the defoaming agent, the microstructure of all geopolymers with graphite was characterized by higher pore volume which is in a good accordance with reduced compressive strength. we can conclude that graphite powder in fly ash geopolymer allows to develop a conductive network within the matrix but also causes generation of additional pores during mixing. the negative impact of graphite powder on the compressive strength is apparently attributed to a weak bond between the matrix and graphite particles because of nonpolar character of graphite surface. low toughness of graphite itself can further contribute to the strength loss at high concentrations. possible ways to reduce these impacts include more efficient defoaming or improved adhesion of graphite particles to the matrix by partial modification of its structure with hydrophilic substituents. acknowledgements this study has been financially supported by the but specific research project fast-j-18-5484 and by ministry of education, youth and sports under the ‘national sustainability programme i’ – the project no. lo1408 admas up. references [1] b. han, l. zhang, j. ou. smart and multifunctional concrete toward sustainable infrastructures, chap. self-sensing concrete, pp. 81–116. springer, singapore, 2017. doi:10.1007/978-981-10-4349-9_6. [2] c. shi, d. roy, p. krivenko. alkali-activated cements and concretes. crc press, 2003. doi:10.1201/9781482266900. [3] v. d. glukhovsky. soil silicates (gruntosilikaty). budivelnik publisher, kiev, 1959. [4] j. davidovits. properties of geopolymer cements. in first international conference on alkaline cements and concretes, pp. 131–149. 1994. [5] l. provis, j. s. van deventer. alkali activated materials. springer, 2014. doi:10.1007/978-94-007-7672-2. [6] p. duxson, a. fernández-jiménez, j. l. provis, et al. geopolymer technology: the current state of the art. journal of materials science 42(9):2917–2933, 2007. doi:10.1007/s10853-006-0637-z. [7] a. fernández-jiménez, i. garcía-lodeiro, a. palomo. durability of alkali-activated fly ash cementitious materials. journal of materials science 42(09):3055– 3065, 2007. doi:10.1007/s10853-006-0584-8. [8] k. p. mehta. reducing the environmental impact of concrete. concrete international 23(10):61–66, 2001. [9] s. hanjitsuwan, s. hunpratub, p. thongbai, et al. effects of naoh concentrations on physical and electrical properties of high calcium fly ash geopolymer paste. cement and concrete composites 45:9–14, 2014. doi:10.1016/j.cemconcomp.2013.09.012. [10] s. hanjitsuwan, p. chindaprasirt, k. pimraksa. electrical conductivity and dielectric property of fly ash geopolymer pastes. international journal of minerals, metallurgy, and materials 18(1):94–99, 2011. doi:10.1007/s12613-011-0406-0. [11] p. payakaniti, s. pinitsoontorn, p. thongbai, et al. electrical conductivity and compressive strength of carbon fiber reinforced fly ash geopolymeric composites. construction and building materials 135:164–176, 2017. doi:10.1016/j.conbuildmat.2016.12.198. [12] s. vaidya, e. n. allouche. strain sensing of carbon fiber reinforced geopolymer concrete. materials and structures 44(8):1467–1475, 2011. doi:10.1617/s11527-011-9711-3. [13] m. saafi, k. andrew, p. l. tang, et al. multifunctional properties of carbon nanotube/fly ash geopolymeric nanocomposites. construction and building materials 49:46–55, 2013. doi:10.1016/j.conbuildmat.2013.08.007. [14] m. saafi, l. tang, j. fung, et al. graphene/fly ash geopolymeric composites as self-sensing structural materials. smart materials and structures 23(6):065006, 2014. doi:10.1088/0964-1726/23/6/065006. 76 http://dx.doi.org/10.1007/978-981-10-4349-9_6 http://dx.doi.org/10.1201/9781482266900 http://dx.doi.org/10.1007/978-94-007-7672-2 http://dx.doi.org/10.1007/s10853-006-0637-z http://dx.doi.org/10.1007/s10853-006-0584-8 http://dx.doi.org/10.1016/j.cemconcomp.2013.09.012 http://dx.doi.org/10.1007/s12613-011-0406-0 http://dx.doi.org/10.1016/j.conbuildmat.2016.12.198 http://dx.doi.org/10.1617/s11527-011-9711-3 http://dx.doi.org/10.1016/j.conbuildmat.2013.08.007 http://dx.doi.org/10.1088/0964-1726/23/6/065006 acta polytechnica ctu proceedings 22:72–76, 2019 1 introduction 2 materials and methods 3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0001 acta polytechnica ctu proceedings 22:1–6, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app the influence of a shrinkage reducing admixture on the long-term development of dynamic properties in alkali-activated slag martin alexaa, dalibor kocába, ∗, barbara kucharczykováa, jan kotrlab a brno university of technology, faculty of civil engineering, veveri 331/95, 602 00 brno, czech republic b brno university of technology, faculty of chemistry, purkynova 464/118, 612 00 brno, czech republic ∗ corresponding author: dalibor.kocab@vutbr.cz abstract. the paper investigates the way a shrinkage-reducing admixture affects the long-term development of dynamic properties in mortars made with alkali-activated slag (aas). two aas mortars were tested one contained a shrinkage-reducing admixture, the other did not. the specimens (prisms of 40 × 40 × 160 mm) were observed for changes in the dynamic modulus of elasticity at ages of 3 days through 2 years using the ultrasonic pulse velocity test and the resonance method. even though the statistical analysis of the results gave no proof that the shrinkage-reducing admixture affected the final values, the development of the elastic modulus appeared to follow a very interesting trend which is completely different from trends commonly observed in cement-based composites. keywords: alkali-activated slag, dynamic modulus of elasticity, ultrasonic pulse velocity test, resonance method. 1. introduction shrinkage, its reduction, or the means of mitigating its consequences are critical points in the development of all types of composite materials. the magnitude of shrinkage depends primarily on the composition of the material and its curing. the type of binder and the progress of hydration play a crucial part as well [1, 2]. alkali-activated binders have been investigated as alternatives to other concrete binders since the 1960s [3]. research shows that that the way alkaliactivated materials perform and behave depends on the raw aluminosilicate material, the activator, and the ratio of the two [4]. the alkali-activated slag composite discussed herein sometimes performs better than composites made with portland cement (pc) – specifically in such properties as durability, resistance to high temperatures, lower hydration heat, or a better interfacial transition zone [5]. the biggest downsides of aas composites are severe shrinkage and rapid setting (especially when the activator is water glass) [6–8]. despite the effort of many investigations to describe the hydration process and mechanical properties of these composites [9–11], some questions pertaining to their behaviour still remain unanswered. shrinkage by desiccation, which is far more severe in aas than in pc [12, 13], results in internal tensile stresses which locally exceed the material’s tensile strength and cause microcracks. this phenomenon can reduce both strength and elasticity. using static tests to observe changes in the structure of the composite and how they may occur over a long period of time is rather time-consuming and requires a large number of specimens. even choosing the right static test is far from straightforward. this is why the structure of these materials is far better observed using dynamic methods based on mechanical wave propagation or forced vibration. these methods can determine the values of the dynamic modulus of elasticity (tensile/compressive, flexural, or shear), which are directly affected by structural changes inside the material. the authors of this paper have already published some of the measurement results (taken between the ages of 3 and 28 days) [14]. however, experience showed that it is better to study the effect of shrinkage-reducing admixtures (sra) over a longer time period. a minimum recommended time is 90 days. 2. experiment 2.1. material and specimens the experiment’s main goal was to determine how a shrinkage-reducing admixture affects the long-term behaviour of aas materials as regards the development of their dynamic modulus. two mortars were tested; both made with granulated blast-furnace slag activated by water glass (sodium silicate). the slag was produced by kotouč štramberk, spol. s r. o. and the water glass (na2o + sio2 + h2o) by vodní sklo, a.s. the water glass was added so that na2o would correspond to 10% of slag mass, resulting in a water-binder ratio of 0.42 (the calculation takes into account water contained in the water glass as well). the aggregate was normalized silica sand cen of a 0–2 mm fraction. the weight of the sand corresponded to three times the weight of slag. the basic components and their ratios were identical in both mortars 1 http://dx.doi.org/10.14311/app.2019.22.0001 http://ojs.cvut.cz/ojs/index.php/app m. alexa, d. kocáb, b. kucharczyková, j. kotrla acta polytechnica ctu proceedings they only differed in the shrinkage-reducing admixture (the commercially available stachement ac600) the first aas contained no admixture and the other contained the sra at 2% of slag weight. all aas components were added to a mixer in the following order: sand, slag, and lastly water with water glass and shrinkage-reducing admixture (if required). the mortars were mixed for 5 minutes at 30 rpm. each of the mortars was made into 9 specimens with the nominal dimensions of 40 × 40 × 160 mm. they were cast in high-strength plastic moulds and stored in a climate chamber with stable ambient conditions – temperature of (21 ± 2) °c and relative humidity of (60 ± 10)% – where they were left to harden without cover or curing for 72 hours. this method was intentional so as to underline the influence of the sra. after the specimens were removed from their moulds, they were left in the same conditions, still without curing, for the next 2 years. 2.2. test methods the experiment focused on observing the young’s modulus e and the shear modulus g. the measurements were made using two non-destructive methods – the ultrasonic pulse velocity test and the resonance method. the basic principle of the former lies in the repeated sending of ultrasonic pulses through the material at very short intervals. a pundit pl-200 with two 150 khz transducers was used to determine the ultrasonic wave transit time, measured longitudinally in three different positions (fig. 1a). the measured data were then used to calculate the ultrasonic wave velocity vl, which is then used for determining the young’s modulus as follows: ecu = ρ · v2l · (1 + µ) · (1 − 2 · µ) 1 − µ , (1) where ecu is the dynamic young’s modulus in mpa, ρ is bulk density in kg/m3, vl is the ultrasonic pulse velocity in km, and µ is the dynamic value of the poisson’s ratio. the dynamic modulus of elasticity was also determined by the resonance method, which operates by introducing vibration to the specimen and detecting the natural frequencies. the specimens received a mechanical pulse delivered by an impact hammer (ih). the natural frequencies were then scanned using a handyscope hs4 oscilloscope with acoustic emission sensors and software that uses a fast fourier transform, see fig. 1b. the natural frequencies were then used to calculate the young’s modulus according to the following formula: ecrl = 4 · l2 · f2l · ρ, (2) where ecrl is the dynamic young’s modulus in mpa, l is the specimen length in m, fl is the measured natural frequency of longitudinal vibration in khz, and ρ is bulk density in kg/m3, or using the formula: ecrl = 0.0789 · c1 · l4 · f2f · ρ · i −2, (3) where ecrf is the dynamic young’s modulus in mpa, c1 is the correction coefficient, l is specimen length in m, ff is the natural frequency of transverse vibration in khz, ρ is bulk density in kg/m3 and i is the radius of gyration of the specimen’s cross-section in m. data obtained by the resonance method can also be used to calculate the shear modulus of elasticity using the formula: gcr = 4 · k · l2 · f2t · ρ, (4) where gcr is the dynamic shear modulus of elasticity in mpa, k is a coefficient determined by the specimen’s cross section, l is specimen length in m, ft is the measured natural frequency of torsional vibration in khz and ρ is bulk density in kg/m3. all nine specimens of each aas mortar were tested for dynamic properties and bulk density at ages of 3, 7, 14, and 28 days. at the age of 28 days, three specimens were tested for the static young’s modulus by compressive loading. the remaining six were later tested for dynamic properties at 90, 180, 365, and 730 days. the pundit pl-200 used the same settings for every measurement – gain: 2× and voltage: 100 v. the resonance method settings were also kept the same throughout the experiment. the dynamic poisson’s ratio used in formula (1) was obtained from the resonance method results. 3. results and discussion fig. 2 through fig. 5 show box plots of the measured dynamic parameters. in order to assess the influence of the shrinkage-reducing admixture, all the data sets, i.e. nine or six values of each property from every measurement date, were verified for normality on a significance level of 0.05. in addition, the test results were statistically compared using a two-sample t-test on a significance level of 0.05. the statistical tests were set up so as to assess the equality of the mean values of the test results. the comparison involved all results for all the ages contrasting aas mortars without shrinkage-reducing admixture (aas-1) with aas mortars that did contain it (aas-2). if the hypothesis is rejected, differences between the sets of results can be considered statistically significant. tab 1 shows the results of the statistical analysis. the data show that none of the mortars reaches values that would be statistically more significant in any of the property being tested at any time during the experiment. an influence of sra on the dynamic moduli of elasticity in aas mortars cannot be identified nor quantified in this case. over the first 14 days of age, the two materials show no statistically significant difference in the values of the young’s modulus, but in other parameters the mortar with no sra has slightly 2 vol. 22/2019 the influence of a shrinkage reducing admixture. . . figure 1. measurement by the ultrasonic pulse velocity test (t transmitter, r receiver); measurement by the resonance method (s acoustic emission sensor, ih impact hammer). age [days] ecu ecrl ecrf gcr 3 aas-1 ≈ aas-2 aas-1 > aas-2 aas-1 > aas-2 aas-1 > aas-2 7 aas-1 ≈ aas-2 aas-1 > aas-2 aas-1 > aas-2 aas-1 > aas-2 14 aas-1 ≈ aas-2 aas-1 > aas-2 aas-1 > aas-2 aas-1 > aas-2 28 aas-1 < aas-2 aas-1 > aas-2 aas-1 > aas-2 aas-1 > aas-2 90 aas-1 > aas-2 aas-1 > aas-2 aas-1 > aas-2 aas-1 ≈ aas-2 180 aas-1 < aas-2 aas-1 < aas-2 aas-1 < aas-2 aas-1 < aas-2 365 aas-1 < aas-2 aas-1 ≈ aas-2 aas-1 < aas-2 aas-1 < aas-2 730 aas-1 > aas-2 aas-1 > aas-2 aas-1 < aas-2 aas-1 < aas-2 table 1. a statistical comparison of test results with and without sra: where the results show no statistically significant difference, the table says “aas-1 ≈ aas-2”; where they do, the table uses a “greater-than” or “less-than” sign accordingly. figure 2. box plot of the development of the young’s modulus ecu. figure 3. box plot of the development of the young’s modulus ecrl. 3 m. alexa, d. kocáb, b. kucharczyková, j. kotrla acta polytechnica ctu proceedings figure 4. box plot of the development of the young’s modulus ecrf . figure 5. box plot of the development of the young’s modulus gcr. higher values. no clear trend can be seen at the age of 28 and 180 days, however, at the age of 1 year the situation becomes opposite to that of 14 days. the mortar with sra performs better at this age; with the exception of the young’s modulus measured by the resonance method (compression wave) where the two mortars do not differ with any statistical significance. at the age of 2 years it again depends which property is being observed – two properties reach significantly higher values in the aas mortar with sra, whereas the remaining two are higher in the mortar without sra. the progress of the aas properties is very interesting, see fig. 6 and fig. 7. they show the progress as a percentage, with 100% being the average 28-day value for all properties. over the first 28 days there is a clear increase in all the properties of both mortars. their young’s modulus reaches over 80% of its 28-day value in as few as 3 days and the shear modulus rises to more than 70%. after 28 days, however, there is a fairly rapid decrease in all the properties. this is where the aas mortar with sra begins to differ from the aas mortar without it. the mortar containing sra saw a drop in the values of the young’s modulus below the 3-day value down to 70%, however, after that they began to slowly rise again – ecu and ecrl return to nearly 90% of their 28-day values in two years, and ecrf and gcr even exceed them. the decrease in all the properties of the mortar without sra continues until the age of 180 days, with their minimum value being lower than in all the properties of the mortar containing the shrinkage-reducing admixture. after that, the mortar also begins to see an improvement; in ecu and ecrl the trend is even more pronounced than in mortar aas-1. 4. conclusion in conclusion, an influence of the shrinkage-reducing admixture on the dynamic moduli of elasticity in aas mortars was not statistically proved. however, the development of the young’s moduli and the shear modulus follows a very interesting trend. over the first 28 days these values increase, followed by a marked decrease, which applies to the mortar without the shrinkage-reducing admixture as well. while the aas mortar with the admixture begins to slowly recover 4 vol. 22/2019 the influence of a shrinkage reducing admixture. . . figure 6. development of the young’s modulus ecu and ecrl, the 28-day value represents 100%. figure 7. development of the young’s modulus ecrf and the shear modulus gcr, the 28-day value represents 100%. at the age of 90 days, the mortar without it continues to suffer a decrease until the age of 180 days when the values begin to rise again. acknowledgements this paper has been written as part of the project no. ga17-14302s “experimental analysis of the early-age volume changes in cement-based composites”, supported by the ga0 – czech science foundation and project no. fast/fch-j-18-5595 “reduction of chemical and autogenous shrinkage of alkaline activated slag”. references [1] h. ye, a. radlińska. shrinkage mitigation strategies in alkali-activated slag. cement and concrete research 101:131–143, 2017. doi:10.1016/j.cemconres.2017.08.025. [2] p. acker. swelling, shrinkage and creep: a mechanical approach to cement hydration. materials and structures 37(4):237–243, 2004. doi:10.1007/bf02480632. [3] j. provis, j. s. j. van deventer (eds.). alkali activated materials: state-of-the-art report, rilem tc 224-aam, vol. 13. 2014. doi:10.1007/978-94-007-7672-2. [4] y. ding, j.-g. dai, c.-j. shi. mechanical properties of alkali-activated concrete: a state-of-the-art review. construction and building materials 127:68–79, 2016. doi:10.1016/j.conbuildmat.2016.09.121. [5] v. bílek, l. kalina, r. novotný, et al. some issues of shrinkage-reducing admixtures application in alkali-activated slag systems. materials 9(6):462, 2016. doi:10.3390/ma9060462. [6] m. palacios, f. puertas. effectiveness of mixing time on hardened properties of waterglass-activated slag pastes and mortars. aci materials journal 108:73–78, 2011. [7] b. kucharczyková, v. bílek jr., d. kocáb, o. karel. shrinkage of fine-grained composites based on alkali-activated slag. in non-traditional cement and concrete, vol. 761 of key engineering materials, pp. 7–10. trans tech publications, 2018. doi:10.4028/www.scientific.net/kem.761.7. [8] v. bílek jr., l. kalina, j. koplík, et al. effect of a combination of fly ash and shrinkage-reducing additives on the properties of alkali-activated slag-based mortars. materiali in tehnologije 50(5):813–817, 2016. doi:10.17222/mit.2015.133. [9] f. collins, j. sanjayan. workability and mechanical properties of alkali activated slag concrete. cement and concrete research 29:455–458, 1999. doi:10.1016/s0008-8846(98)00236-1. [10] c. shi, r. l. day. a calorimetric study of early hydration of alkali-slag cements. cement and concrete research 25(6):1333–1346, 1995. doi:10.1016/0008-8846(95)00126-w. [11] f. collins, j. sanjayan. cracking tendency of alkali-activated slag concrete subjected to restrained shrinkage. cement and concrete research 30:791–798, 2000. doi:10.1016/s0008-8846(00)00243-x. 5 http://dx.doi.org/10.1016/j.cemconres.2017.08.025 http://dx.doi.org/10.1007/bf02480632 http://dx.doi.org/10.1007/978-94-007-7672-2 http://dx.doi.org/10.1016/j.conbuildmat.2016.09.121 http://dx.doi.org/10.3390/ma9060462 http://dx.doi.org/10.4028/www.scientific.net/kem.761.7 http://dx.doi.org/10.17222/mit.2015.133 http://dx.doi.org/10.1016/s0008-8846(98)00236-1 http://dx.doi.org/10.1016/0008-8846(95)00126-w http://dx.doi.org/10.1016/s0008-8846(00)00243-x m. alexa, d. kocáb, b. kucharczyková, j. kotrla acta polytechnica ctu proceedings [12] a. neto, m. cincotto, w. repette. drying and autogenous shrinkage of pastes and mortars with activated slag cement. cement and concrete research 38(04):565–574, 2008. doi:10.1016/j.cemconres.2007.11.002. [13] f. collins, j. sanjayan. effect of pore size distribution on drying shrinking of alkali-activated slag concrete. cement and concrete research 30(09):1401– 1406, 2000. doi:10.1016/s0008-8846(00)00327-6. [14] d. kocáb, l. topolář, v. bílek jr., et al. experimental analysis of the development of compressive strength, modulus of elasticity and acoustic emission parameters of alkali-activated composites. in special concrete and composites 2017, vol. 760 of key engineering materials, pp. 266–271. trans tech publications, 2018. doi:10.4028/www.scientific.net/kem.760.266. 6 http://dx.doi.org/10.1016/j.cemconres.2007.11.002 http://dx.doi.org/10.1016/s0008-8846(00)00327-6 http://dx.doi.org/10.4028/www.scientific.net/kem.760.266 acta polytechnica ctu proceedings 22:1–6, 2018 1 introduction 2 experiment 2.1 material and specimens 2.2 test methods 3 results and discussion 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.16.0006 acta polytechnica ctu proceedings 16:6–10, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app prediction of the development of groundwater quality in a finished sludge lagoon petr černocha, ∗, jiří košťálb a čez energetické produkty, s. r. o., komenského 534, 253 01 hostivice, czech republic b inset, s. r. o., divize energetika, lucemburská 7, 130 00 praha 3, czech republic ∗ corresponding author: cernoch.petr@inset.com abstract. this study deals with prediction of the development of groundwater quality in a finished sludge lagoon. for depositing and predicting secondary float energy by-products, it is appropriate to use mathematical modelling. the implementation of relevant input data from long-term monitoring is suitable as an application for verification as well. the main output was the determination of the necessary priorities and geotechnical risks in solving the final state of the area of interest, especially regarding solving of the drainage system in the subsequently finished sludge lagoon. keywords: transport of contaminants, sludge bed, geotechnical risks. 1. introduction when depositing secondary energy by-products from coal-fired power plants, several crucial questions arise in relation to the ambient environment and potential geotechnical risks. regarding the remediation and revitalisation of former sludge beds by way of depositing secondary energy by-product, it is necessary to point out fundamental phenomena related to the impact of the structure of interest on the ambient environment and potential geotechnical risks. this concerns mainly a rise in groundwater level due to superimposed load on the structure [1]. when the secondary energy by-products are deposited, this new layer can sag and consequently, open tensile fractures may appear in it. therefore, it is advisable to pay close attention to the analysis of changes that can arise when former sludge beds are covered with a remediation layer. in particular, there is a risk of restoring the problem as the groundwater outflow conditions may change, which can lead to the transport of potential contaminants from the former sludge beds to the ambient environment. at the same time, the stability of some parts of the structure may be affected locally and the overall stability in a wider area may also be affected, especially if the uplift pressure at the base of the structure increases [2]. the above-mentioned risks may be disclosed in time and subsequently eliminated thanks to the implementation of hydrogeological model calculations which focus on the simulation of changes in outflow conditions caused by depositing secondary energy by-products. typical tasks mainly include monitoring the impact of hydro-regime changes on the stability of a former sludge bed (or rather the finished sludge lagoon), verification of possible spreading of potential contaminants in the saturated ground environment, and groundwater level development in various alternative solutions for overall capacity extension and reclamation of the structure [3, 4]. the prediction of groundwater quality development (modelling of contaminant transport) was implemented in the area of two (former) sludge beds. the first sludge bed was created in a depressed terrain by enclosing it with a starter dam. it was hydraulically operated between 1967 and 1990. the neighbouring sludge bed was located in the area of a former quarry in which mining was discontinued in 1983. ash deposition by way of sluicing in the central part of the structure of interest, started in 1990 and has continued to this day while the sluicing is scheduled to stop at the end of 2017 or the beginning of 2018. a hydromixture (a mixture of slag and water) is sluiced to the sludge bed. simultaneously, the implementation since 1997 of coal-fired power plant desulphurisation using the wet lime scrubbing process, a pre-reclamation arrangement for landscape recovery has been modelled on the surface of a sludge bed ash beach using a stabilised mixture prepared from secondary energy by-products (a mixture of fly ash, lime, wastewater, etc.). 2. input data first, a geological model was created and hydrodynamic (hydraulic) properties assigned to partial geological units. after that, boundary conditions were defined and then a computing grid was established. for the partial tasks to be solved, it was essential to first solve the hydrogeological model with steady groundwater flow in a wider area of interest. thus, the precondition of setting actual initial and boundary conditions at the border of the model area is met. visual modflow flex 2015.1. was used for the modelling, including the implemented modules: modflow 2005 u.s. geological survey to find out the course of the groundwater flow in the saturated ground 6 http://dx.doi.org/10.14311/app.2018.16.0006 http://ojs.cvut.cz/ojs/index.php/app vol. 16/2018 ground water quality in sludge lagoon environment, mt3mds for the transport of contaminants and zonebudget to establish the groundwater amount in pre-defined 3d bodies (areas). afterwards, vmod 3d-explorer enabled us to display both inputs and outputs of individual models in 3d form. 2.1. model geometry and boundary conditions the locality of interest was modelled over a total surface area of approx. 3.2×2.8 km (i.e., an area of approx. 905 hectares was simulated). the lower limit of the model was selected in such a way that the total vertical profile adequately covered all the geological units concerned and their general course. great emphasis was laid on creating geological boundaries of tuffitic rocks (forming an impermeable barrier between the two sludge beds) and coal, including coalbearing clays (especially the course of the coal seam and its non-balanced residues). the geological model was identical for all partial numerical models, only the input of the terrain surface course differing in the variants. the current terrain, the filled body after slag sluicing was discontinued (12/2017) and after that depositing of stabilised mixture is finished, that is before the reclamation starts in 2025. when the geological model was implemented (figure 1), the conceptual model was solved, and later there followed the numerical model. the conceptualisation consisted in the schematisation of all aspects of the area of interest, i.e., the geometry of the analysed area and the groundwater flow conditions. moreover, the model material composition was simplified while respecting all the defined boundary conditions. a specific constant groundwater level was used as the first-type boundary condition for the modelled borders of the area of interest. the level was not entered as constant up to the border but as respecting the found data (groundwater level) from the hydrogeological model of the wider surroundings of the area of interest. in addition, a second-type boundary condition was introduced, namely recharge. to avoid separate entry of another boundary condition – evapotranspiration (difficult determination of amount and depth range), was already included when the recharge boundary condition was entered – and this value was reduced. superficial saturation and local watercourses were included in the model as a 3rd-type boundary condition – river. another 3rd-type boundary condition was the course of levels in the stream tubing (drain) and drainage system for the sludge bed seepage water. the boundary condition of the lake included the artificially maintained constant height of the free water level in the residual sludge bed lake and hydro-mixture sluicing. water recharge from the sluiced hydro-mixture can be derived from the known value of the hydro-mixture flow rate through the sluice pipeline minus the slag content. the entered 4th-type boundary condition, pumping wells, represented the effect of overflow towers. the initial condition (initial head) with the starting groundwater level calculated within the steady groundwater flow modelling in the wider area of interest, was used in the entire simulated model area. 2.2. material characteristics the following basic geological units were selected for the solved models: stabilised mixture; slag-ash sediment; stockpiled soil; holocene deposits including deluvial-fluvial soils; aeolian or aeolian-deluvial sediments of loess or of loess-soil nature; fluvial pleistocene gravel accumulations – terraces; overlying part of most formation – clays with sand fraction; coal series (seam and non-balanced residues) – coal; coal claystone; pyroclastic rock including the underlying part of most formation – tuffs and tuffits of volcanic series; varied claystones and tuffitic clays of underlying series; cretaceous rock of turonian age – marlites and lime claystones; vulcanites (neogene effusive magmatites) – olivine basalts, nephelinites and leucites. the selection of all 11 various geological environments made for a certain simplification consisting in the unification of soils and rocks into so-called quasi-homogeneous units (i.e., units with similar geological and hydrogeological properties), although they may be heterogeneous as to their lithology and genesis. hydrodynamic parameters for each geological unit were obtained from hydrodynamic tests, well logging, laboratory analyses, and also partially derived from archive searches in czech and foreign professional literature [5]. the determination of the ash hydraulic parameters is very difficult. from experience in other similar fields, it is known that filtration coefficients in a saturated ash body in analogous conditions, as verified by a pumping experiment, showed differences of one or two orders of magnitude. differences between filtration coefficients in the horizontal and vertical directions may even be a hundredfold. moreover, in the central part of the sludge bed (near the overflow towers) the finest-grain ash fraction tends to settle and, mainly thanks to the existence of a free water surface in this area, the difference between the vertical and horizontal permeability values is almost entirely blurred. for those reasons, flatly introduced filtration coefficient values were adequately adjusted within the modelling. at the same time, crevice permeability was not considered in the solved model calculations. it plays a part in underlying cretaceous rocks and the stabilised mixture body. considering the position of the cretaceous deposits (deeper groundwater circulation), only a negligible error results in the calculation. it is true for both the slag-ash sediment and stabilised mixture that in reality (on the macro-scale), their permeability will be greater than the laboratory value. however, this is insignificant in the presented calculations. within the solution, the reference groundwater levels obtained from long-term monitoring of the area of interest were entered. additionally, earlier chemical 7 petr černoch, jiří košťál acta polytechnica ctu proceedings figure 1. initial geological model (superelevated 10 fold). analyses were searched out, groundwater samples were taken, and their subsequent laboratory analyses undertaken. the input values entered in the calculation model were average values obtained from the reference period from 2009 to 2016. the groundwater level values measured in the monitoring hydro-wells served for the calibration and backward verification of the model. the latter compared whether the calculated course of the groundwater level in the model initial time corresponded to the actually found levels to the maximum extent, i.e., the long-term average in recent years. 3. calculation, calibration and backward verification when solving the transport of contaminants in the area of interest, it was necessary to implement partial groundwater flow models. in general, four basic models were solved: • model 1: groundwater amount and quality in the area of the sludge beds in simultaneous slag sluicing to the other sludge bed; • model 2: prediction model of groundwater quality after the sluicing is discontinued for selected chemical indicators (until 2025); • model 3: prediction model of the groundwater amount flowing out of the area of both sludge beds from the discontinuation of slag sluicing to the other sludge bed till the discontinuation of depositing secondary energy by-product – final shape before reclamation (until 2025); • model 4: prediction model of the groundwater level when the water regime has stabilised after the discontinuation of sluicing. except for the first and partly also the last model, with the steady state of groundwater, a so-called “steady-state simulation”, the variants were always time-varying. the groundwater course reacting to time changes was solved, as a response to the change in the entered boundary conditions. in the case of model 1, we used the standard computing module of pcg2 (preconditioned conjugate-gradient method) to numerically solve the simulated equations of the linear groundwater flow model using modified incomplete cholesky and polynomial predictions. for the other models, the computing method of gmg (geometric multigrid solver) served the purpose of solving the corresponding matrices. this method is based on the conjugate gradient algorithm like the previous pcg2, but unlike pcg2 it was developed especially for solving groundwater flow models with a numerical grid of finite differences. for models 1 and 2 (after timevariant groundwater flow iterations were successfully finished) the advective flow method of ufd (upstream finite difference) with the implicit computing module of gcg (generalized conjugate gradient solver) was used to calculate the transport of contaminants. once the obtained results had been reviewed, we proceeded to the calibration and backward verification of the numerical model. hydrogeological model calibration consisted in a gradual change in the input parameters until an acceptable agreement between the measured and calculated values was achieved. in the first step, the achieved groundwater levels in zero time were compared with the measured data in each model variant. then the concentration values for the monitored chemical indicators were calibrated as well. numerical model verification was made possible mainly thanks to a sufficient amount of data obtained 8 vol. 16/2018 ground water quality in sludge lagoon from the long-term monitoring of the area of interest performed for many years by inset s.r.o. in summary, we can conclude that, despite its extreme complexity (extensive territory, complicated geological structure, large number of boundary conditions, etc.), the implemented model of the solved area achieved, in all its partial variants, a very good agreement with the documented values from the hydrogeological and hydrochemical monitoring. 4. results of model calculations all model variants included the effect of the sludge bed drainage system. in the implemented mathematical modelling, the effect of the upper near-surface aquifer was connected with deeper groundwater circulation. thus, the resulting groundwater level courses always reflected the shallowest continuous aquifer in the locality of interest, which comprised the saturated zone of both sludge beds, the quaternary aquifer at the inflow (sandy gravels), the near-surface aquifer of the shallow groundwater circulation connected in the east to a broad alluvial plain of the original stream bed (connecting to surface waters), and also in the north-east part of the second sludge bed a partial show of tertiary water recharge (coal seam outbursts). thus, the determined groundwater level represented the highest possible level of the continuous course. considerable deep groundwater circulation (mostly a crevice aquifer in cretaceous sediments and tertiary vulcanites and deposits) was found under the modelled course of the shallow groundwater level (pore character of the aquifer) and had virtually no bearing on the transport of contaminants solved by us; therefore, it was not included in the implemented simulations. the modelled groundwater level course imitated to a large extent the terrain sloping and slope direction. furthermore, it was obvious that the closer to the residual lake the calculated groundwater level was, the more the shallow groundwater levels gradually evened out, approaching the level corresponding to the free water surface of the sludge lagoon. therefore, the simulated steady groundwater flow indicates communication between the groundwaters from both sludge beds. in the case of model 1, the groundwater quality was shown as constant, representing the current state. the comparison of calculated values of the monitored chemical indicators with average values documented by monitoring in the reference period from 2009 to 2016 showed very good agreement. the achieved values of individual potential contamination parameters served as input data for the further solution variant implemented in model 2. model 2 solved the groundwater quality prediction for the following chemical indicators: dissolved inorganic salts (ras), ammonium ions, sulphates and arsenic. as mentioned above, it was based on the current state calculation for groundwater quality (model 1) and simultaneously on model 3, which focused on the change in outflow conditions, or rather the state of the groundwater levels from sluicing discontinuation till the creation of the final secondary energy by-product body, i.e., till 2025. the simulation of groundwater quality development from the beginning of secondary energy by-product depositing is quite a complicated operation; therefore, it is partially burdened with error. from the initial phase of the numerical calculation, i.e., from the time when the first sludge bed operation started in 1967, a significant leap in groundwater quality can be observed from this “zero” point up to the current state. it follows that the gradual deterioration of contamination up to the current state (for approx. 50 years) has a long-term tendency and therefore, no rapid or considerable improvement can be expected in the upcoming years (i.e., from the time when sluicing / secondary energy by-product depositing is discontinued). the calculated values of the monitored parameters had a relatively good agreement with the measured values of these chemical indicators in the long term. inaccuracies in simulations compared with the measured data got worse towards the borders of both sludge beds. the mathematical calculation of the current state was followed by the calculation of groundwater quality prediction in time, i.e., its change from 12/2017 till 2025. it was obvious in all the monitored cases that the equilibrium of the hydrochemical environment did not hold steady at the time when secondary energy by-product depositing was discontinued. the contaminant cloud moved very slowly in time and the contaminants will most probably never be washed completely out of both sludge beds, because of the lack of a sufficient gradient and (ground) water recharge, as shown in the results of model 4. it should be noted that according to the calculations, the water regime will stabilise in approximately 2042, which means 25 years after sluicing to the sludge bed is discontinued and 17 years after depositing of the stabilised mixture is discontinued (before reclamation). in the case of all chemical indicators, the modelling showed that all their monitored values decreased in time after sluicing was discontinued. at the end of the solved period (from approximately 2022), a partial decrease in the monitored concentration parameters started to show also in wells located at the outflow from both sludge beds. a detectable decrease in values of the monitored concentration parameters inside the sludge bed bodies would require a longer time period. for most indicators, the concentration values documented by the model for the final year 2025 were lower in all monitored wells than the maximum value found in the selected monitoring period of 2009 – 2016. in half of the cases, the calculated values (for 2025) were lower than the average values in the above-mentioned reference period. 9 petr černoch, jiří košťál acta polytechnica ctu proceedings 5. conclusions we can conclude that the indicated development of groundwater quality showed a clear tendency. that is, after discontinuation of sluicing, the contaminant cloud was gradually washed out, diluted and moving outside the sludge bed. however, the development depended not only on depositing of the secondary energy by-product but primarily on the hydraulic transport operation. if, theoretically, the sludge bed was supplied only with clear water through the sluice pipeline, the contaminant cloud could be gradually pushed out of the sludge bed and almost eliminated in the end. it means the cloud movement also implies a certain dilution, i.e., a decrease in contaminant concentrations. it is also quite obvious that the washing-out of contaminants will be considerably affected by the volume of inflows from the gravel aquifer. to the contrary, if substantial recharge of the sluicing water is missing, both concentration decrease and contaminant cloud movement will slow down; according to numerical modelling there will be an increase in concentrations of the monitored parameters inside both sludge bed bodies. another issue connected with the potential adverse effect of finishing the sludge lagoon (a change in hydroregime and thus the stability of the area of interest) is maintaining the drainage system function until the water regime completely stabilises. the performed model calculations have implied that it is necessary to operate the drainage system at least until the groundwater flow stabilises, i.e., for approx. 15 years after depositing of the secondary energy by-product in the area of interest is discontinued. none of the model cases required any adjustment of limit values in the number of internal and mainly external iterations in order to increase the tolerance for achieving convergence above the size of one-tenth of a computing cell. a very good agreement with initial times was achieved for the values obtained from the model and the field measurement, thanks to calibration and backward verification. the main uncertainty in the results obtained was the issue of course continuity of the groundwater level. the numerical calculation did not allow modelling of isolated saturated places; it is intrinsically made to solve a continuous course of the groundwater level. this results in a certain simplification and distortion; in reality, there will be separated, preferential groundwater flows in some isolated localities which will also allow the transport of contaminants. references [1] p. černoch, j. košťál. studie vlivu změn hydrorežimu na stabilitu odkališť. in proceedings of conference geotechnics 2016. orgware, 2016. [2] p. černoch, j. košťál. implementation of geotechnical risks during realization of a slag dam based on the sediments in a former sludge lagoon. in proceedings of 16th conference ecsmge. ice publishing, 2015. [3] p. černoch, j. košťál. řešení ukládání vedlejších energeticých produktů do prostoru odkaliště. in proceedings of 14th congress of hydrogeology and 2nd congress of engineering geology. čah, 2014. [4] p. černoch, j. košťál. stanovení zatěsnění úložného prostoru vedlejších energetických produktu na základě experimentu in situ. geotechnika journal 2, 2016. [5] p. černoch, j. košťál. uplatnění hydrogeologického modelování při ukládání vep. in proceedings of 3rd international conference popílky ve stavebnictví and 27th international symposium sanace 2017. vut brno, 2017. 10 acta polytechnica ctu proceedings 16:6–10, 2018 1 introduction 2 input data 2.1 model geometry and boundary conditions 2.2 material characteristics 3 calculation, calibration and backward verification 4 results of model calculations 5 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0115 acta polytechnica ctu proceedings 13:115–120, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of failure criteria for wood in tensile-shear stress state eliška šmídová∗, petr kabele czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: eliska.smidova@fsv.cvut.cz abstract. an orthotropic failure criterion enhancing the lourenco's criterion by a shear strength multiplier and a maximum shear strength upper bound has been recently proposed and validated for timber under tensile and shear loading by the authors. the paper discusses its applicability for predicting strength in comparison with tsai-hill criterion, hankinson's and hyperbolic formula applying the two above mentioned enhancements of the lourenco's criterion. experimental data available in the literature for off-axis tensile and shear test of sitka spruce (picea sitchensis carr.), katsura (cercidiphyllurn japonicurn sieb. and zucc.), douglas fir (pseudotsuga menziesii), douglas fir laminated veneer and cupiúba (goupia glabra) are used for the purpose of this study. keywords: orthotropic failure criterion, timber, off-axis tensile test, shear strength upper bound. 1. introduction a large number of phenomenological strength criteria, that apply to the failure phenomenon in general instead of the failure mechanism itself, has been proposed for composite materials in recent decades. similarly to composites, wood is typical for twodimensional character and organized non-isotropic structure and as such can be categorised as orthotropic or transversely isotropic. it follows that strength theories developed for artificial composites can be used for prediction of wood strength [1, 2]. unfortunately, for some of these theories, it may be difficult to obtain specific coefficients experimentally, let us mention the issue of the interaction term [3–5]. all strength parameters may not be viewed in a deterministic way. stochastic analysis is an alternative approach that is on the rise. it takes into consideration, among others, correlation of strength properties that we can treat as independent random variables [2]. the off-axis test has become a widely used experiment for validation of failure criteria of non-isotropic materials. at a relatively low cost and operational requirements, it provides the strength of tensile or compressive specimens under uniaxial loading at various load-grain angles. specimens with the angle of 0°and 90°yield tensile strength parallel and perpendicular to the grain, respectively. the off-axis tensile test has been involved in numerous studies, especially in those related to the complex problem of determining shear strength of wood [6–10] or the applicability of failure criteria for a specific wood species [11, 12]. in the present study, we compared four existing failure criteria: (i) lourenco’s criterion modified for wood by the authors [13, 14], (ii) tsai-hill criterion [15], (iii) hankinson's formula [16] and (iv) hyperbolic formula [17]. furthermore, we applied some of the modifications of the lourenco's criterion to the other criteria. we made the comparison against the data available in the literature for sitka spruce, katsura [7], douglas fir [18], douglas fir laminated veneer [19] and cupiúba [20]. 2. failure criteria in this section, we give a brief overview of four failure criteria that are involved in this study. the formulas are expressed in terms of either a 2d stress state or the off-axis test using the following relations: σx = σ0m2 (1) σy = σ0n2 (2) τxy = −σ0nm (3) 2.1. lourenco's criterion modified for timber failure criterion that has been recently proposed by the authors [13, 14] is defined by means of rankine failure surface modified for orthotropy [21]. it is further enhanced by (i) parameter ps that multiplies shear strength fs in the failure surface formula, (ii) shear strength as an upper bound of the shear stress and (iii) crack type criterion. the first two enhancements enable a calibration of the failure surface through offaxis test results. crack type criterion defining whether a crack is initiated either along or across the grain is out of the scope of this work. failure surface and failure criterion are expressed in a 2d stress space by equation (4) and (5), respectively. 115 http://dx.doi.org/10.14311/app.2017.13.0115 http://ojs.cvut.cz/ojs/index.php/app eliška šmídová, petr kabele acta polytechnica ctu proceedings τxy,fs =   psfs ftxfty √ ftxfty(σx − ftx)(σy − fty) for τxy < fs fs for τxy ≥ fs (4) τxy = τxy,fs (5) 2.2. tsai-hill criterion the most frequently used failure theory for wood was proposed by norris [15]. it is also referred to as the tsai-hill criterion. it is an extension of the von miseshencky distortion energy hypothesis. in a 2d stress state, it has the form of equation (6). ( σx ftx )2 − σxσy ftxfty + ( σy fty )2 + ( τxy fs )2 = 1 (6) 2.3. hankinson’s formula this formula [16] has been used mostly for predicting the ultimate compressive strength of wood. some researchers have applied it also for predicting tensile strength [22]. the formula is as follows: σ0 = ftxfty ftxnh + ftymh . (7) 2.4. hyperbolic formula the hyperbolic formula has been developed to provide a better fit to douglas fir off-axis test results compared to other commonly used failure theories [17]. the definition is following: σ0 = 2ftxfty exp0.01θ(ftx + fty) + exp−0.01θ(fty − ftx) . (8) 3. results and discussion using the equations 1 3, the failure criteria are expressed in terms of the off-axis tensile test, i.e. σ0(θ), and compared with the experimental data from the literature for off-axis tensile and shear tests. for the purpose of calibration, two enhancements of the lourenco's failure criterion proposed by the authors (section 2.1) are applied to the other failure conditions, if possible. these are the use of: (1.) a shear strength multiplier ps in the failure surface formula (psfs instead of fs), (2.) shear strength as an upper bound of shear stress in a 2d stress space. 3.1. sitka spruce and katsura yoshihara and ohta conducted off-axis tensile and shear test for sitka spruce (picea sitchensis carr.) and katsura (cercidiphyllurn japonicurn sieb. and zucc.) [7]. specimens were conditioned at 65% relative humidity. for the first experiment type, dog-bone specimens with the outer dimensions of 140×10×8 mm were cut, five pieces for each grain angle out of 0°, 5°, 10°, 15°, 20°, 25°, 30°, 45°, 60°, 75° and 90°. shear strength was obtained from both off-axis tensile test (by transformation of the axial strength to the stress components of orthotropic symmetry) and torsion test. for the latter, ten dog-bone specimens with the outer dimensions of 180×20×5 mm were prepared. in this experimental campaign, shear strength from torsion test coincided well with that from the off-axis test for 15°– 30°. off-axis tensile test and shear test results for sitka spruce together with the uncalibrated failure curves (ps = 1.0, h = 1.0) are shown in the upper part of figure 1 in terms of off-axis tensile test variables. we can see that out of the uncalibrated curves, the tsai-hill curve represents a good estimate. calibrating the modified lourenco's, tsai-hill and hankinson's curves by the parameters of ps = 2.0, ps = 1.2 and h = 1.8, respectively, and activating the shear strength upper bound, we obtain the best fit of the average experimental data, see the lower part of figure 1. on the contrary, the hyperbolic curve does not reproduce the data well. experimental data for katsura are plotted together with uncalibrated and calibrated failure curves in the upper and lower part of figure 2, respectively. similarly to the results for sitka spruce, tsai-hill curve is the best estimate out the uncalibrated curves. nevertheless, by the calibration of the modified lourenco's, tsai-hill and hankinson's curves by the parameters of ps = 2.3, ps = 1.3 and h = 2.2 , respectively, and by application of the shear strength limit to the shear stress, we get a good prediction of the data. the hyperbolic curve strongly underestimates the data. 3.2. douglas fir woodward and minor measured tensile strength parallel and perpendicular to fibers and shear strength for douglas fir (pseudotsuga menziesii) [18] according to the specimen configuration and testing procedure specified by astm d-143 [23]. they used eight specimens for each type of experiment. for determination of the ultimate strength at grain angles of 15°, 30°, 45° and 60°, they used eight rectangular specimens for each grain angle. the specimens were 38 mm wide, 10 mm thick and they were of different length, excluding the gripping area: 480 mm, 330 mm, 250 mm and 200 mm for the grain angles of 15°, 30°, 45° and 60°, respectively. the specimens were dried to the moisture content of 12 ± 1%. uncalibrated and calibrated failure curves are shown in the upper and lower part of figure 3, respectively, together with the experimental results for douglas fir. both the modified lourenco's and hyperbolic criterion fit the average off-axis tensile test data well without any calibration. on the other hand, we can reproduce the data even better applying ps = 0.8 116 vol. 13/2017 comparison of failure criteria for wood 3 [°] 0 10 20 30 40 50 60 70 80 90 < 0 [ m p a ] 0 20 40 60 80 100 120 140 160 modified lourenco p s = 1.0 modified tsai-hill p s = 1.0 hankinson h = 1.0 hyperbolic off-axis avg off-axis avg'cov f s,avg f s,avg'cov 3 [°] 0 10 20 30 40 50 60 70 80 90 < 0 [ m p a ] 0 20 40 60 80 100 120 140 160 modified lourenco p s = 2.0 modified tsai-hill p s = 1.2 hankinson h = 1.8 hyperbolic off-axis avg off-axis avg'cov f s,avg f s,avg'cov figure 1. the experimental data of sitka spruce [7] with uncalibrated (top) and calibrated (bottom) failure conditions. and h = 1.5 to the tsai-hill and hankinson's formulas. let us note that for this wood species, the shear strength upper bound is not activated. in this study, we disregard size effect following conclusions of a related experimental work [24]. 3.3. douglas fir laminated veneer clouston et al. performed off-axis tensile test and shear test for douglas fir laminated veneer [19]. offaxis specimens of equal size (610×63×35 mm) were cut with grain angles of 0°, 15°, 30°, 45°, 60° and 90°, sixteen to eighteen pieces for each angle. shear strength was evaluated on nineteen standard astm shear block specimens using shear adjustment factor [25]. average moisture content of the specimens was 7.9%. looking at uncalibrated failure curves and experimental data in the upper part of figure 4, we can see that the modified lourenco's curve represents the best estimate. we can get similar results if we calibrate the tsai-hill's and hankinson's criteria by ps = 0.65 and h = 1.8, respectively, see the lower part of figure 4. we can notice that the shear strength upper bound is not activated. 3 [°] 0 10 20 30 40 50 60 70 80 90 < 0 [ m p a ] 0 20 40 60 80 100 modified lourenco p s = 1.0 modified tsai-hill p s = 1.0 hankinson h = 1.0 hyperbolic off-axis avg off-axis avg'cov f s,avg f s,avg'cov 3 [°] 0 10 20 30 40 50 60 70 80 90 < 0 [ m p a ] 0 20 40 60 80 100 modified lourenco p s = 2.3 modified tsai-hill p s = 1.3 hankinson h = 2.2 hyperbolic off-axis avg off-axis avg'cov f s,avg f s,avg'cov figure 2. the experimental data of katsura [7] with uncalibrated (top) and calibrated (bottom) failure conditions. 3.4. cupiúba todeschini conducted off-axis tensile and shear test for cupiúba (goupia glabra) [20] in accordance with the brasilian norm nbr 7190 [26]. for the purpose of the off-axis test, dog-bone specimens with the outer dimensions of 280×20×20 mm were cut, twelve pieces for each grain angle out of 0°, 15°, 30°, 45°, 60°, 75° and 90°. shear strength was measured using shear block specimens. the moisture content of the specimens varied from 12% to 14%. plotting the uncalibrated curves together with the experimental results of cupiúba (the upper part of figure 5), we can see that the hyperbolic curve yields the closest estimate of the average off-axis data. the lower part of figure 5 shows that we can reproduce well the data utilizing ps = 0.6, ps = 0.4 and h = 1.7 for the modified lourenco's, tsai-hill and the hankinson's criteria, respectively. similarly to the case of douglas fir and douglas fir laminated veneer, the shear strength upper bound is not activated. 4. conclusion the paper compares four orthotropic failure criteria against off-axis and shear experimental results avail117 eliška šmídová, petr kabele acta polytechnica ctu proceedings 3 [°] 0 20 40 60 80 100 < 0 [ m p a ] 0 20 40 60 80 100 120 140 modified lourenco p s = 1.0 modified tsai-hill p s = 1.0 hankinson h = 1.0 hyperbolic off-axis avg off-axis avg'cov f s,avg f s,avg'cov 3 [°] 0 20 40 60 80 100 < 0 [ m p a ] 0 20 40 60 80 100 120 140 modified lourenco p s = 1.0 modified tsai-hill p s = 0.8 hankinson h = 1.5 hyperbolic off-axis avg off-axis avg'cov f s,avg f s,avg'cov figure 3. the experimental data of douglas fir [18] with uncalibrated (top) and calibrated (bottom) failure conditions. able in the literature for the wood species of sitka spruce, katsura, douglas fir, douglas fir laminated veneer and cupiúba. the enhancements of the modified lourenco's criterion, which has been recently proposed and validated by the authors, are applied to the tsai-hill, hankinson's and hyperbolic formula, to a certain extent. specifically, (i) the shear strength multiplier and (ii) the shear strength upper bound are used. following conclusions can be drawn: (1.) the shear strength multiplier enables calibration of a failure formula that contains shear strength. in this way, both, the modified lourenco's and tsaihill criteria provide a better estimation of the average off-axis data for sitka spruce, katsura, douglas fir, douglas fir laminated veneer and cupiúba. (2.) the shear strength upper bound, that can be activated as a maximum limit of the shear stress component, provide a better fit for average off-axis data of sitka spruce and katsura. for these wood species, average shear strength measured by torsion test coincides well with the average off-axis strength. (3.) calibration of the hankinson's formula by the parameter h yields good estimates for all wood species involved in this study, similarly as the modified 3 [°] 0 10 20 30 40 50 60 70 80 90 < 0 [ m p a ] 0 10 20 30 40 50 60 70 modified lourenco p s = 1.0 modified tsai-hill p s = 0.8 hankinson h = 1.5 hyperbolic off-axis avg off-axis avg'cov f s,avg f s,avg'cov 3 [°] 0 10 20 30 40 50 60 70 80 90 < 0 [ m p a ] 0 10 20 30 40 50 60 70 modified lourenco p s = 1.0 modified tsai-hill p s = 0.65 hankinson h = 1.8 hyperbolic off-axis avg off-axis avg'cov f s,avg f s,avg'cov figure 4. the experimental data of douglas fir laminated veneer [19] with uncalibrated (top) and calibrated (bottom) failure conditions. lourenco's and tsai-hill criteria do. in contrast, the hyperbolic formula predicts well only douglas fir data as it does not contain any calibration or shear strength parameter. list of symbols θ load-grain angle [°] ftx tensile strength parallel with grain [mpa] fty tensile strength perpendicular to grain [mpa] fs shear strength [mpa] h parameter of the hankinson's criterion [–] m cos θ [–] n sin θ [–] ps shear strength multiplier [–] σ0 ultimate strength at a load-grain angle θ [mpa] σx normal stress parallel with grain [mpa] σy normal stress perpendicular to grain [mpa] τxy shear stress [mpa] acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs17/168/ohk1/3t/11. 118 vol. 13/2017 comparison of failure criteria for wood 3 [°] 0 10 20 30 40 50 60 70 80 90 < 0 [ m p a ] 0 20 40 60 80 100 modified lourenco p s = 1.0 modified tsai-hill p s = 1.0 hankinson h = 1.0 hyperbolic off-axis avg off-axis avg'cov f s,avg 3 [°] 0 10 20 30 40 50 60 70 80 90 < 0 [ m p a ] 0 20 40 60 80 100 modified lourenco p s = 0.6 modified tsai-hill p s = 0.4 hankinson h = 1.7 hyperbolic off-axis avg off-axis avg'cov f s,avg figure 5. the experimental data of cupiúba [20] with uncalibrated (top) and calibrated (bottom) failure conditions. references [1] r. s. sandhu. a survey of failure theories of isotropic and anisotropic materials. tech. rep., air force flight dynamics lab wright-patterson afb oh, 1972. [2] b. kasal, r. j. leichti. state of the art in multiaxial phenomenological failure criteria for wood members. progress in structural engineering and materials 7(1):3–13, 2005. doi:10.1002/pse.185. [3] s. w. tsai, e. m. wu. a general theory of strength for anisotropic materials. journal of composite materials 5(1):58–80, 1971. [4] a. a. zolochevskii. tensor relationship in the theories of elasticity and plasticity of anisotropic composite materials with different tensile and compressive strength. mechanics of composite materials 21(1):41–46, 1985. [5] r. e. rowlands, d. e. gunderson, j. c. suhling, m. w. johnson. biaxial strength of paperboard predicted by hill-type theories. the journal of strain analysis for engineering design 20(2):121–127, 1985. doi:10.1243/03093247v202121. [6] e. p. saliklis, r. h. falk. correlating off-axis tension tests to shear modulus of wood-based panels. journal of structural engineering 126(5):621–625, 2000. [7] h. yoshihara, m. ohta. estimation of the shear strength of wood by uniaxial-tension tests of off-axis specimens. journal of wood science 46(2):159–163, 2000. [8] j. y. liu. analysis of off-axis tension test of wood specimens. wood and fiber science 34(2):205–211, 2002. [9] h. yoshihara, t. satoh. examination of the off-axis tension test method for evaluating the shear properties of wood. forest product journal 53(5):75–79, 2003. [10] j. c. xavier, n. m. garrido, m. oliveira, et al. a comparison between the iosipescu and off-axis shear test methods for the characterization of pinus pinaster ait. composites part a: applied science and manufacturing 35:827–840, 2004. [11] n. t. mascia, r. todeschini, e. a. nicolas. uniaxial and biaxial test for evaluation of failure criterion of anisotropic material applied to wood. in world conference on timber engineering, riva del garda, italy. 2010. [12] n. t. mascia, r. a. simoni. analysis of failure criteria applied to wood. engineering failure analysis 35:703–712, 2013. doi:10.1016/j.engfailanal.2013.07.001. [13] e. bartunkova. constitutive model of timber. master’s thesis, czech technical university in prague, 2013. [14] e. smidova, p. kabele. constitutive model for timber fracture under tensile and shear loads. applied mechanics and materials 784:137–144, 2015. doi:10.4028/www.scientific.net/amm.784.137. [15] c. b. norris. strength of orthotropic materials subjected to combined stresses, 1962. [16] r. l. hankinson. investigation of crushing strength of spruce at varying angles of grain. air service information circular 259, 1921. [17] j. woodward, b. clinton. the elastic and strength properties of douglas fir in the radial-longitudinal plane. ph.d. thesis, new mexico state university, new mexico, 1986. [18] c. woodward, j. minor. failure theories for douglas-fir in tension. journal of structural engineering 114(12):2808–2813, 1988. [19] b. p. clouston, f. lam, j. d. barrett. interaction term of tsai-wu theory for laminated veneer. journal of materials in civil engineering pp. 112–116, 1998. [20] r. todeschini. ensaios uniaxiais e biaxiais para avalicao de criterio de resistencia de materiais anisotropicos aplicado a madeira. master’s thesis, faculdade de engenharia civil, arquitetura e urbanismo, campino, brasil, 2009. [21] p. b. lourenco, r. de borst, j. g. rots. a plane stress softening plasticity model for orthotropic materials. int j numer methods eng 40:4033–4057, 1997. [22] f. f. p. kollmann, w. a. cote. principles of wood science and technology: i solid wood. springer berlin heidelberg, berlin, heidelberg, 1968. oclc: 851389657. [23] a. s. d-143. standard methods for testing small clear specimens of timber, vol. 09. american society for testing and materials, philadelphia, pa., 1984. 119 http://dx.doi.org/10.1002/pse.185 http://dx.doi.org/10.1243/03093247v202121 http://dx.doi.org/10.1016/j.engfailanal.2013.07.001 http://dx.doi.org/10.4028/www.scientific.net/amm.784.137 eliška šmídová, petr kabele acta polytechnica ctu proceedings [24] t. tannert, t. vallee, s. franke, p. quenneville. comparison of test methods to determine weibull parameters for wood. world 15:19, 2012. [25] p. clouston. the tsai-wu strength theory for douglas-fir laminated veneer. master’s thesis, the university of british columbia, vancouver, 1995. [26] abnt. nbr 7190: projeto de estruturas de madeira. abnt, rio de janeiro, 1996. 120 acta polytechnica ctu proceedings 13:116–121, 2017 1 introduction 2 failure criteria 2.1 lourencos criterion modified for timber 2.2 tsai-hill criterion 2.3 hankinson's formula 2.4 hyperbolic formula 3 results and discussion 3.1 sitka spruce and katsura 3.2 douglas fir 3.3 douglas fir laminated veneer 3.4 cupiúba 4 conclusion list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.20.0029 acta polytechnica ctu proceedings 20:29–33, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app calibration of btt measurement with respect to sensor position over shrouded lsb tereza dadáková∗, zdeněk kubín doosan škoda power s.r.o., tylova 1/57, pilsen, czech republic ∗ corresponding author: tereza.dadakova@doosan.com abstract. in order to increase efficiency and reliability of steam turbines, last stage blades are equipped with mid-span tie-boss and shroud. we can use blade tip-timing method (btt) to measure vibration of shrouded blades, but the measured level of vibration is dependent on position of sensor over blade shroud. hence, the results of btt measurement have to be compared with other measurement techniques, like strain gauges, and with fem analysis, to interpret these results right. this article provides findings we made on measurement of shrouded blades in test rig with eddy-current sensors, optical sensors and strain gauges at different positions of sensors over blade shroud. measurement is also compared to fem analysis. keywords: blade tip-timing, calibration, sensor. 1. introduction blade tip timing (btt) measurement is in doosan skoda power used mainly for two reasons: to prevent blade failure on some operational states, where can be increased level of vibration because of ventilation unpredictable asynchronous vibration can be excited by small aerodynamic forces. for these cases, longterm monitoring with calculation of blade lifetime is used on power plants to plan shutdowns of turbine and measuring blade behaviour on wide range of operation points to optimalize operation of the turbine. second use of btt in doosan skoda power is blade testing in campbell test rig to compare measurements with fem analysis and prove that the blades are designed out of synchronous resonances. for shrouded blades we use campbell test rig to do the calibration of btt determination of suitable position for sensor location, comparison of results of btt with strain gages and verification of calculation. the need to do the calibration of btt occurred with the new design of blades with interconnecting elements. to last stage blade (lsb) design tie-bos in the middle of the blade and shroud at the tip of the blade were added. the centrifugal force causes untwisting of blades, resulting in force between these contact elements and hence in reduction of blade vibration. however these features brought many questions about measuring those blades with btt. the blade shroud has a great impact on signal shape from the sensor, in some positions (contact plane) it is complicated to receive a good data. another question is what point of shroud is exactly measured and it is necessary to interpret the amplitude of vibration for btt right. also the identification of cracks on shrouded blades became more demanding task. cracks usually cause a change in blade vibration behaviour like frequency shift, higher level of vibration, shift of a blade towards other blades change in inner blade distance, etc. however on free blades is this behavior manifested more significantly. an alternative measurement to btt are contact methods use of strain gages. strain gages give us direct information about the blade stress, but usually are installed only on few blades and in corrosive steam enviroment with high temperature usually have a short service life. also the installation is quite expensive and these reasons make this method unsuitable for wider use. btt as contactless method, or nsms (nonintrusive stress measurement system), is cheaper, sensors are placed circumferentially around stator and hence each sensor sees each blade of the rotating disk [1]. the btt system measures time arrival of each blade, the difference between time arrivals is converted to vibration in mm. this data can be analyzed in many ways, using various methods [2, 3]. most used method is fast fourier transformation (fft) converting the signal from a sensor to single-blade spectrums or allblade spectrum. single-blade spectrum gives us information about vibration of every blade. however, it is sampled only once per revolution what causes limited frequency range and problems with identification of natural frequency and nodal diameter. on the other side, all-blade spectrum shows us the whole disk vibration behaviour with possibility of determination nodal diameters and natural frequencies (it is sampled by number of blades in one revolution), but changes in vibration of one blade with defect may not occur in this spectrum. this usually leads to using a combination of different methods to receive a real amplitude of all blades. various types of sensors can be used, new sensors with different properties are developed [4]. sensors can be based on different physical principals, not in every case it is clear what exactly sensors measure. simulation of measurement can help here [5, 6]. 29 http://dx.doi.org/10.14311/app.2018.20.0029 http://ojs.cvut.cz/ojs/index.php/app tereza dadáková, zdeněk kubín acta polytechnica ctu proceedings this article contains observations based on measurement of shrouded blade in campbell test rig, using optical sensor, eddy-current sensors and strain gages. this measurement was special, because of using a system that gradually moved sensors over rotating blades. this way we were able to “scan” blade shroud and find best sensor position for measuring this type of blade. it also gave us a different point of view on measuring with eddy-current sensors. we also describe a method we use to calibrate btt with calculation. 2. measurement description the measurement was done in campbell test rig on lsb for 60 hz with diameter 2904 mm, the disk consisted of 50 blades. the campbell test rig used for measurement operates without steam, blades rotate in vacuum, because high-speed motion through air would result in aerodynamic heating and rise of temperature. the level of vacuum is usually about 3 mbar, maximal rotational speed is 4000 rpm and maximal disk diameter is 5 m. to excite blades an electro-magnet, oil jets or torsional shaking can be used. the electo-magnet excites blades on frequency fe, fe = fn + nd · fr, (1) where fn is natural blade frequency, nd is nodal diameter which is excited and fr is rotational frequency. electro-motor torsional shaking is used to excite 0 nd, because from our experience the amplitude is ten-times bigger, than by using the electromagnet. unfortunately, only this nodal diameter can be excited this way. 2.1. measuring system to measure the blade vibration contact and contactless methods were used, as mentioned in section 1. in the test rig it is possible to use strain gages with telemetry system, btt and high speed camera monitoring system. for this measurement only strain gages and btt were used. strain gages were placed at positions, where we expected maximal response to verify shapes of various nodal diameters. amplitudes are obtained from fast fourier transformation directly on natural frequency. btt consists of sensor, amplifier, trigger and measuring system with high sampling frequency clock (figure 1). in the test rig we successfully use optical sensors, but in steam enviroment the use of optical sensors is problematic. for use in steam enviroment eddycurrent sensors are more suitable, because they are not affected by steam and are designed to operate in temperatures up to 250 ◦c. this leads to the need of comparison of those two types of sensors, since we know, what exactly optical sensors measure, but what eddy-current measure exactly on shrouded blade is a question. eddy-current design is shown in the figure 2. the signal from eddy-current is sensitive on position of magnet and coil inside the sensor and on size of shroud compared to the size of sensor. the shape of signal is very dependent on position of sensor over blade shroud and on rotational frequency. the signal from sensor is amplified (a in figure 1) and transferred to ttl digital puls (trigger t in figure 1). the measuring system with clock (s in figure 1) measures time arrivals of edges of ttl pulses. the system accuracy is dependent on sampling frequency of measuring clock. for example, if the rotational frequency fr is 60 hz and rotor diameter d = 2904 mm, to see vibration of order 10−1 mm we would need sampling frequency about fs = 5.5 mhz, to see vibration of order 10−2 mm frequency fs = 55 mhz and to see vibration of order 10−3 mm even fs = 550 mhz. longer blades need even higher sampling frequency, disk with diameter 5 m would have blade tip speed about 940 m/s, so to measure with 10−3 mm accuracy, frequency fs = 940 mhz would be needed. for this measurement the sampling frequency was 200 mhz, which provided accuracy about 3 µm. 2.2. measurement setup we used optical sensor and two similar eddy-current sensors with slightly different properties (size, number of loops in coil). those sensors were attached to a threaded rod with stepper motor, that allowed us to move those sensors along the blade shroud step by step to measure on different positions without requirement of running-down and moving the sensors manually (figure 3). we were able to chose the speed of sweep and position arbitrarily and we saved values of the positions so we could compare it with the measurement. by the optical sensor we measured both, leading and trailing edge of the blade (figure 4), so we were able to capture the whole blade shroud shape. results are summarized in next sections. 3. results we have done two types of measurement: first, we investigated the influence of magnet and coil position to blade and secondly we excited blades on its natural frequencies to compare amplitudes from different sensors. we didn’t move the sensors continuously, but we stopped every one mm for a few seconds to capture the vibration on every position. 3.1. sensor orientation we performed five measurements, where we measured the blade shroud with sensor moving there and back starting on trailing edge, so we were sure, that the captured shape is not affected by level of vibration or inaccurate movement of sensors. we chose sensor orientation as shown in figure 5, where circle marks the coil in the sensor. starting by positioning the coil towards the blade, next positions were determined by turning the sensor by 90◦, 180◦, 210◦ and 300◦. 30 vol. 20/2018 calibration of btt measurement a t s 1/rev so se figure 1. diagram of btt measuring system. coil magnet figure 2. eddy-current sensor. electro-magnet optical sensor eddy-current sensor stepper motors figure 3. measurement setup. figure 4. edges captured by optical sensor. to ilustrate the result we chose optical sensor and one eddy-current. those sensors centers were distant from each other 190 mm (1.04 blade distance) and the data form each sensor are related to the 1/rev sensor. if we choose one specific blade, for example the first one that arrives to optical sensor after 1/rev mark, it is actually the second one, that arrives to eddy-current. the position of k-th blade of sensor s in relative angle pk in revolution n can be expressed as pk,n = toask,n − toarn toarn+1 − toarn , (2) where toa stands for time of arrival, r for sensor 1/rev and s for optical sensor, resp. eddy-current. we take k-th blade from optical sensor and k + 1-th blade from eddy-current, recalculate relative angle to mm, subtract the distance of sensors and the result is shown in figure 6. data from eddy-currents are displayed as colored dots, colors correspond to orientation of sensor from figure 5. the shape of blade shroud is taken from the optical sensor data. te in the figure 6 stands for trailing edge and le for leading edge. the sensor size is for this blade comparable with the size of the blade shroud, what is in our opinion the cause of some uncertainty of what eddy-current measures, especially on leading edge. from experience it is expected, that the sensor measures by the center of the coil, which unfortunately wasn’t confirmed by this measurement. this measurement was very important, because it confirmed, that on trailing edge is the 31 tereza dadáková, zdeněk kubín acta polytechnica ctu proceedings figure 5. eddy-current sensor orientation. le te figure 6. eddy-current sensor data (relative angle position of one blade in disk.) measurement almost independent on orientation and that here it captures the blade shroud shape appropriately (even if it may not be exactly on the edge). trailing edge is just the position, that is mainly used for btt measurement. more in section 4. 3.2. amplitude on natural frequency another measurement we performed, was exciting the bladed disk on its natural frequency and comparison of amplitude of optical and eddy-current sensors. we excited the blades on frequency fe, see equation 1, where the exact natural frequency fn was determined from strain gages measurement. we again measured the blade shroud there and back with few seconds stops, so we were able to use fft on data in one position with sufficient accuracy. to compare amplitude form both sensors we used amplitude from all-blade spectrum. from parts of the data, where the sensors were steady, the all-blade fft was computed. the amplitude on excitation frequency fe from this spectrum was taken and that is represented by color in the figure 7. in the figure 7 are results for two natural frequencies, 0nd and 6nd. as mentioned in section 2, we use torsional shaking for excitation of 0nd and electro-magnet for higher nodal diameters. from the figure 7 0nd results we can see, that on trailing edge both sensors measure so so so so se se so so so so se se figure 7. amplitude for optical and eddy-current sensors on 0nd (top) and 6nd (bottom). so optical sensor, se eddy-current. the same amplitude (color of eddy-current and optical sensor matches). that shows us the optimal position for measuring this blade. unfortunately, as we can see in results for 6nd, eddy-current sensor measurement was affected by the electro-magnet excitation increasingly, as the sensor was getting closer to the electro-magnet, that was placed on the leading edge side of the blade, see figure 3, where its position is shown. for higher nodal diameters only data from optical sensor from this measurement were reliable. 32 vol. 20/2018 calibration of btt measurement yc figure 8. deflection measured by btt. yc y1 y2 y1 α β figure 9. recalculation of deflection measured by btt to the actual deflection. 4. calibration method let us conclude with a description of how we use the results we get from this measurement and how we compare it with calculation. the btt sensor measures deflection in the direction of rotation, in the figures 8 and 9 marked as yc, however the actual vibration of blade in this direction is y1. using angles α and β we have yc = ( 1 tan α tan β + 1 ) y1. (3) from here comes the importance of angles α and β. the angle α is given by blade design and angle β by mode shape is different for every nodal diameter and is computed by fem. if we take measurement from strain gages in µm/m and divide it by actual amplitude measured by btt, we get a ratio that expresses strain on one mm of blade tip deflection for specific strain gage position. this can be also calculated using fem and those two values we compare to verify computed blade properties. 5. conclusion the measurement brought a new look on measuring with eddy-current sensor and we found an optimal position of sensor over blade shroud for this type of blade. the results may be extended to graphical comparison of measured data to fem analysis depending on nodal diameter, type of sensor etc. to get better results with excitation higher nodal diameters using eddy-current sensors it is necessary to find a way of exciting the blades without affecting the sensor signal. we expect a repetition of this measurement in the test rig with another type of blade with larger diameter. this could demonstrate the difference of sensor signal behaviour, especially on leading edge, where the measuring point of eddy-current for this blade is unclear. list of symbols fs sampeling frequency of measuring clock [hz] fe excitation frequency [hz] fn natural frequency [hz] fr rotational frequency [hz] n d nodal diameter pk,n relative angle of k-th blade in n-th revolution t oask,n time of arrival of k-th blade in n-th revolution to sensor s t oarn time of arrival of 1/rev marker in n-th revolution yc displacement seen by btt [mm] y1 actual displacement of blade in tangential direction [mm] acknowledgements these research’s results are part of the turbo-reflex project, which has received funding from the european union’s horizon 2020 research and innovation programme under grant agreement no 764545. references [1] z. kubin, j. liska, t. misek, p. prochazka. application of non-contact vibration measurements on a 48” blade under low load condition. dymamesi 2017 prague: institute of thermomechanics as cr pp. 45–54, 2017. [2] g. battiato, c. firrone, t. berruti. forced response of rotating bladed disks: blade tip-timing measurements. mechanical systems and signal processing 85(1):912–926, 2017. doi:10.1016/j.ymssp.2016.09.019. [3] o. danek, f. vanek, j. kozanek, et al. the development of identification of dynamic parametres of rotating machinery parts. proceedings of the conf structural system identification 2001. [4] p. prochazka, f. vanek. new methods of non-contact sensing of blade vibrations and deflections in turbomachinery. ieee trans instrum meas 63:1583–1592, 2014. [5] n. jamia, m. friswell, s. el-borgi, r. fernandes. simulating eddy current sensor outputs for blade tip timing. advances in mechanical engineering 10(1), 2018. [6] z. kubin, t. misek, j. hlous, et al. calibration of blade tip-timing sensor for shrouded 40” last stage blade. mechanical systems and signal processing 108:88–98, 2018. doi:10.1016/j.ymssp.2018.02.001. 33 http://dx.doi.org/10.1016/j.ymssp.2016.09.019 http://dx.doi.org/10.1016/j.ymssp.2018.02.001 acta polytechnica ctu proceedings 20:29–33, 2018 1 introduction 2 measurement description 2.1 measuring system 2.2 measurement setup 3 results 3.1 sensor orientation 3.2 amplitude on natural frequency 4 calibration method 5 conclusion list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0109 acta polytechnica ctu proceedings 15:109–113, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app constitutive model for timber fracture used for fe simulation of lvl arch eliška šmídová∗, petr kabele czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: eliska.smidova@fsv.cvut.cz abstract. non-linear finite element simulation of four-point bending test of a small-size arch is described in this paper. the arch of 780 mm span is made of yellow poplar (liriodendron tulipifera) laminated veneer lumber (lvl) and it is manufactured with a through crack parallel to fibers in the middle of its crown. 2d homogeneous orthotropic constitutive model of tensile and shear fracture in timber that has been recently developed and implemented into atena® finite element software by the authors is used for the numerical calculation. both standard and compact tension (ct) test results are used for the material model calibration. results show that the model successfully reproduces the increasing part of the load-displacement response. furthermore, the model can capture the most distinctive features of the crack pattern. keywords: timber, 2d constitutive model, orthotropic elasticity, orthotropic tsai-hill failure criterion, fixed smeared crack model, traction-separation law, compact tension test. 1. introduction concerning environmental impact, energy efficiency and handling costs, timber and engineered wood products represent a promising alternative to the most frequently used structural materials such as concrete or masonry. the wood-derived products of laminated veneer lumber (lvl), cross-laminated timber (clt) or glue-laminated timber (glt) provide increased strengths and stiffness compared to solid wood. timber and its engineered products are typical for orthotropic behavior. sawn timber has three mutually perpendicular axes in which we distinguish independent mechanical properties: longitudinal (parallel to grain), radial and tangential. let us note that properties, such as tensile strength and modulus of elasticity, related to the direction parallel with fibers substantially exceed those related to the other directions [1]. natural variability in material properties of timber has been addressed by some authors [2, 3]. this material feature is often tackled by weibull strength theory [4]. exposed to tensile and shear loading, timber exhibits linear elastic behavior up to failure that is followed by brittle fracture. in contrast, compression in timber leads to a ductile response [5]. various failure criteria have been developed for timber [6, 7] out of which a commonly applied one corresponds to a tsai-hill formulation [8]. being a complex biological structure, timber contains natural imperfections such as knots, pitch pockets, reaction wood or cross grain. because such inhomogeneities have a negative impact on the mechanical properties, structural timber, especially the engineered one, is usually carefully selected and processed to provide an improved structural behavior [1]. influence of some of these weaknesses (e.g., distribution of knots and finger joints) to strength and failure pattern of lvl beams is tackled in [9], among others. two-dimensional and three-dimensional phenomenological constitutive models capturing timber behavior under multi-axial stress state are based on elasticity [10], classical plasticity [11], multisurface plasticity [12], smeared crack concept [5, 13], smeared fictitious reinforcing fibers [14] and interface schemes [15–17]. 2. constitutive model for non-linear finite element simulation of 4pbt of yellow poplar (liriodendron tulipifera) laminated veneer lumber (lvl) arch, we use the 2d homogeneous orthotropic constitutive model of tensile and shear fracture in timber modified for available input data. the model has been recently developed and implemented into atena® finite element software [18] by the authors [13, 19]. 2.1. model properties let us review the properties of the modified constitutive model. it captures: • material orthotropy for small deformations in both linear and non-linear range • elastic behavior until the anisotropic failure criterion is fulfilled • inelastic behavior after failure criterion is satisfied • post-failure response for cracks along the grain • unloading/reloading behavior the post-cracking response is treated by the fixed smeared crack model [20] where reinhardt-hordijk function is used [21]. 109 http://dx.doi.org/10.14311/app.2018.15.0109 http://ojs.cvut.cz/ojs/index.php/app eliška šmídová, petr kabele acta polytechnica ctu proceedings 2.2. model components the properties of the model described in section 2.1 are implemented in the form of (i) failure criterion and (ii) cohesive law. 2.2.1. failure criterion with regard to available input data from literature, we use the most frequently used failure theory for wood. it was proposed by norris [8] and it is also referred to as the tsai-hill criterion. in a 2d stress state, it has the form of equation eq. 1. ( σx ftx )2 − σxσy ftxfty + ( σy fty )2 + ( τxy fs )2 = 1 (1) 2.2.2. cohesive law cohesive (traction-separation) law relates crack surface normal traction tn to relative crack normal displacement δn with respect to normal crack direction (n). the cohesive law is expressed by eq. 2 as follows: tn(δn) =   fty (( 1 + ( c1δn δn,crit )3) exp ( c2δn δn,crit ) − exp (−c2) ( 1 + c31 ) δn δn,crit ) for δn ∈ [0,δn,crit) 0 for δn > δn,crit (2) 3. calibration of the model the model requires in total ten input parameters. based on data available in the literature, we obtained material input parameters for yellow poplar lvl both directly from standard tests and employing inverse analysis of traction-separation relationship from compact tension (ct) test (figure 1). representing first part of input parameters for fe simulation of lvl arch, selected standard tensile and shear experimental results [1, 17] are shown in table 1. within experimental test campaign of yellow poplar lvl arch available in literature [22], only two load vs. crack mouth opening displacement (cmod) responses of ct specimens cut from the arch base and three load vs. displacement responses of ct specimens cut from the arch crown were reported. it is important to note, that displacement of the latter curves is not relevant being biased by the stiffness of the load-cell. thus, we estimated the parameters defining cohesive law in crack-normal direction, i.e. δn,crit, fty, c1, c2, in two steps. first, for each of two responses of ct specimens from the arch base, we get the average of four tractionseparation curves using sequentially linear approach to inverse analysis including multi-pass enhancement [23]. consequently, we find the best fit of the cohesive law eq. 2 for the average curve. similarly, we find the best fit for the average of all eight traction-separation curves. the overview of cohesive law parameters figure 1. compact tension (ct) test set up: w = 46 mm, a = 20 mm and thickness is 11 mm. ey ex ftx fs gxy νxy [mpa] [mpa] [mpa] [mpa] [mpa] [-] 450 14600 110 8.2 899 0.32 table 1. average values of input data for yellow poplar lvl from standard tests [1, 17]. calculated for ct specimens from the arch base is shown in table 2 together with r2 statistics that witness a good fit of corresponding responses from finite element simulation and experiment. second, for each of three values of maximum load from response of ct specimens sawn from the arch crown, we estimate tensile strength parameter fty considering remaining parameters of {c1,c2,δn,crit} to be equal to those calculated in step one for ct specimens from arch base. for this purpose, we apply the trial-and-error method to fe inverse analysis. a considerable difference between maximum loads applied to ct specimens cut from the arch base and crown can be observed. therefore, we use the estimations of cohesive law parameters calculated for the arch crown (table 3) expecting them to be more representative for fe simulation of the crack pattern spreading across the crown. the responses (load vs. cmod) and traction-separation relationships (stress vs. cod) from (i) experiments, (ii) sequentially linear or (iii) trial-and-error approach to inverse analysis are displayed in figure 2 and figure 3, respectively. 4. fe simulation of 4pbt of the arch the subject of the fe simulation is a four-point bending test (4pbt) of yellow poplar lvl arch with a through pre-crack parallel with grain in the middle of its crown [22, 24]. teflon paper was placed between two laminations preventing them from being glued. the arch has the span of 780 mm, cross-section height and width of 57 mm and 25 mm, respectively, and lamina thickness of 3.2 mm. the overall test setup is depicted in figure 4. we generated regular 2d fe mesh composed of approx. 3.2 mm by 3.2 mm quadrilaterals aligned with 110 vol. 15/2018 constitutive model for timber fracture response δn,crit fty c1 c2 r 2 [mm] [mpa] [-] [-] [-] max 1.41 15.90 4.10 7.90 0.93 avg 1.07 14.00 4.10 8.60 − min 0.57 14.00 4.10 10.00 0.96 table 2. parameters of cohesive law calculated by sequentially linear approach to inverse analysis [23] of load vs. cmod responses of ct specimens cut from arch base. {δn,crit,c1,c2} fmaxty f avg ty f min ty {[mm], [−], [−]} [mpa] [mpa] [mpa] {1.41, 4.10, 7.90} 2.60(#1) 2.10(#2) 1.60 {1.07, 4.10, 8.60} 2.90(#3) 2.30(#4) 1.80 {0.57, 4.10, 10.00} 3.90 3.00 2.30 table 3. parameters of cohesive law estimated by trial-and-error approach to inverse analysis of max. load responses of ct specimens cut from arch crown. the arch outline and lamina. we assigned fiber orientation to each fe as an input material property. as the proposed material model treats post-cracking behavior by the smeared crack approach, we modeled the pre-crack as a missing row of fe elements. let us note, that in this way we get the same results as for the pre-crack modeled as a row of fes with an unloaded smeared crack at the beginning of a numerical simulation. we ran fe simulations of lvl arch for each of 9 estimates of input parameters defining the traction-separation relationship for the arch crown table 3. 5. results and discussion we present four most relevant results are presented in figure 5 in the form of load vs. midspan deflection. out of them, we calculated the best response for sim. (#1) (i.e., maximum estimated critical crack opening displacement δn,crit = 1.41 mm and tensile strength fmaxty = 2.6 mpa). regarding significant variability of timber material parameters, the results show that the model can reproduce the increasing part of the load-displacement response well. in order to obtain statistically more relevant results, an experimental campaign containing numerous specimens for each type of test needs to be conducted. comparison of crack pattern across arch crown from experiment (figure 6) and fe simulation (figure 7) shows that the material model can capture both the crack propagation from the pre-crack tips (primary crack) and crack initiation of the crack below the precrack (secondary crack). however, the model neglects initiation of the crack above the pre-crack (tertiary crack). the short crack spacing of only twice the lamina thickness between primary and tertiary crack can cause that stored energy is released by opening figure 2. load vs. cmod responses of ct specimens cut from arch base and crown: (i) experimental (exp.), (ii) numerically simulated by means of sequentially linear approach to inverse analysis (sim.) and (iii) estimated by trial-and-error approach to inverse analysis (estim.). and spreading of the primary crack instead of the adjacent tertiary crack initiation and growth. in the end, we can compare the results for yellow poplar lvl arch from smeared crack model to the results calculated by interface scheme [22, 24]. to this end, let us briefly describe the latter approach first. interface elements were placed between the lamina and the initial crack was modeled as a gap of finite length separating the arch layers. interface behavior was governed by (i) modulus of elasticity and (ii) maximum tensile strength, both in the direction perpendicular to fibers, and (iii) fracture toughness acquired from compact tension specimens cut from arch crown [16]. the response obtained by the interface scheme is shown in figure 5 as sim. interf. this curve represents a weak reproduction of the response compared to sim. (#1) to sim. (#4). on the other hand, the interface approach can capture both the primary crack spreading and the secondary crack initiation and growth. discrepancies between the two approaches can be attributed to: (1.) significant variability of timber material parameters that needs to be handled appropriately. (2.) specific acquisition of the most relevant input parameters that control the crack initiation and post-cracking behavior. 111 eliška šmídová, petr kabele acta polytechnica ctu proceedings figure 3. traction-separation relationships (stress vs. cod) of ct specimens cut from arch base and crown: (i) experimental (exp.), (ii) numerically simulated by means of sequentially linear approach to inverse analysis (sim.) and (iii) estimated by trial-and-error approach to inverse analysis (estim.). figure 4. scheme of 4pbt of yellow poplar lvl arch with a through pre-crack in its crown. 6. conclusion we used 2d homogeneous orthotropic constitutive model of tensile and shear fracture in timber for fe simulation of small-size yellow poplar lvl arch subjected to four-point bending. we presented the main features of the model. input data were obtained both (i) directly from standard tensile and shear tests and (ii) by applying inverse analysis to compact tension (ct) test results for yellow poplar laminated veneer lumber (lvl). for the limited availability of relevant experimental data from ct test results, we calculated the estimation of parameters that define traction-separation relation. numerical simulations of yellow poplar lvl arch confirm that the applied constitutive model can capture both the increasing part of the load-displacement response and pre-crack spreading together with the initiation of a secondary crack. we compared the results obtained by utilizing the proposed model to those calculated by interface approach. both models can predict primary and secondary crack initiation figure 5. load vs. midspan deflection response of 4pbt of lvl arch obtained from experiments (exp.) and fe simulations (sim.). i i.i. ii. iii. figure 6. arch crown cracking (primary i., secondary ii. and tertiary iii.) during experiment at two load levels. and development well. on the other hand, the interface model yields a weak load vs. midspan deflection response. considerable variability of timber material characteristics requires both statistically significant results for each type of experiment and an appropriate statistical approach to obtain relevant results from numerical models. list of symbols c1,c2 material parameters controlling the shape of the softening curve [–] δn crack opening displacement [mm] δn,crit critical value of crack opening displacement [mm] ex modulus of elasticity along the grain [mpa] 112 vol. 15/2018 constitutive model for timber fracture figure 7. arch crown cracking in fe simulation: primary and secondary crack. ey modulus of elasticity across the grain [mpa] ftx tensile strength parallel with grain [mpa] fty tensile strength perpendicular to grain [mpa] fs shear strength [mpa] gxy shear modulus [mpa] νxy poisson’s ratio [–] r2 coefficient of determination [–] σx normal stress parallel with grain [mpa] σy normal stress perpendicular to grain [mpa] τxy shear stress [mpa] acknowledgements the support by the gačr grant no. 18-05791s is gratefully acknowledged. references [1] wood handbook wood as an engineering material. forest product laboratory, 2010. 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[16] x.-p. xu, a. needleman. numerical simulations of fast crack growth in brittle solids. journal of the mechanics and physics of solids 42(9):1397–1434, 1994. doi:10.1016/0022-5096(94)90003-5. [17] b. kasal, r. blass. experimental and analytical investigation of crack development in composite reinforced laminated arch. materials and structures 46(1-2):173–180, 2013. doi:10.1617/s11527-012-9892-4. [18] j. cervenka, l. jendele, v. cervenka. atena program documentation, part 1, theory, 2008. [19] e. bartunkova. constitutive model of timber. ph.d. thesis, czech technical university in prague, 2013. [20] a. hillerborg, m. modeãľr, p.-e. petersson. analysis of crack formation and crack growth in concrete by means of fracture mechanics and finite elements. cement and concrete research 6(6):773–782, 1976. [21] h. reinhardt, h. cornelissen, d. hordijk. tensile tests and failure analysis of concrete. journal of structural engineering 112(11):2462–2477, 1986. [22] r. a. blass. laminated wooden arches reinforced with glass fiber rods. ph.d. thesis, the pennsylvania state university, 2011. [23] j. vorel, p. kabele. inverse analysis of traction-separation relationship based on sequentially linear approach. arxiv:180208040 [cs] 2018. arxiv: 1802.08040. [24] b. kasal, r. blass. experimental and analytical investigation of crack development in composite reinforced laminated arch. materials and structures 46(1-2):173–180, 2013. doi:10.1617/s11527-012-9892-4. 113 http://dx.doi.org/10.4028/www.scientific.net/amm.611.54 http://dx.doi.org/10.1002/pse.185 http://dx.doi.org/10.14311/app.2017.13.0115 http://dx.doi.org/10.1016/j.engstruct.2012.08.017 http://dx.doi.org/10.1007/s10086-012-1296-0 http://dx.doi.org/10.1007/s00466-003-0423-6 http://dx.doi.org/10.4028/www.scientific.net/amm.784.137 http://dx.doi.org/10.1016/j.conbuildmat.2011.03.017 http://dx.doi.org/10.4028/www.scientific.net/ssp.35-36.287 http://dx.doi.org/10.1016/0022-5096(94)90003-5 http://dx.doi.org/10.1617/s11527-012-9892-4 http://dx.doi.org/10.1617/s11527-012-9892-4 acta polytechnica ctu proceedings 15:109–113, 2018 1 introduction 2 constitutive model 2.1 model properties 2.2 model components 2.2.1 failure criterion 2.2.2 cohesive law 3 calibration of the model 4 fe simulation of 4pbt of the arch 5 results and discussion 6 conclusion list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2016.6.0034 acta polytechnica ctu proceedings 6:34–39, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app the dubins traveling salesman problem with constrained collecting maneuvers petr váňa∗, jan faigl dept. of computer science, czech technical university in prague, technická 2, 166 27 prague, czech republic ∗ corresponding author: vanapet1@fel.cvut.cz abstract. in this paper, we introduce a variant of the dubins traveling salesman problem (dtsp) that is called the dubins traveling salesman problem with constrained collecting maneuvers (dtsp-cm). in contrast to the ordinary formulation of the dtsp, in the proposed dtsp-cm, the vehicle is requested to visit each target by specified collecting maneuver to accomplish the mission. the proposed problem formulation is motivated by scenarios with unmanned aerial vehicles where particular maneuvers are necessary for accomplishing the mission, such as object dropping or data collection with sensor sensitive to changes in vehicle heading. we consider existing methods for the dtsp and propose its modifications to use these methods to address a variant of the introduced dtsp-cm, where the collecting maneuvers are constrained to straight line segments. keywords: dubins vehicle, dtsp, non-holonomic vehicle, path planning, collecting maneuver. 1. introduction the dubins traveling salesman problem (dtsp) [1] has been subject of studies for many years. the problem is to find a shortest tour visiting a given set of target locations by the dubins vehicle [2] that models a fix-wing aircraft (or car-like vehicle) with a limited minimal turning radius and a constant forward velocity. the herein studied path planning problem is motivated by surveillance and rescue missions to visit a set of target locations by a fix-wing aerial vehicle. such a problem can be formulated as the variant of the combinatorial optimization the traveling salesman problem (tsp) where the target locations have to be connected by dubins maneuvers. the problem is to determine a sequence of visits to the targets together with the most suitable heading of the vehicle at each target location such that the total tour length is minimized. dubins proved that the optimal path between two configurations of the dubins vehicle can be found analytically by enumerating six possible maneuvers [2]. however, it is necessary to determine the optimal headings to find the optimal maneuver analytically, which in fact depend on the optimal sequence of visits to the targets. therefore, this challenging problem is called the dubins traveling salesman problem (dtsp) rather than just the tsp to distinguish computational challenges arising from the constrained turning radius. the dtsp is np-hard [3] as for zero minimal turning radius, the problem becomes the tsp. even though an ordinary formulation of the dtsp provides a solution of the problem to visit a set of target locations by an aircraft, the problem considered in this paper requires to perform specific collecting maneuvers to reliably accomplish the mission. such a requirement is arising from the need to get a visual contact with the drop-off location. then, the vehicle figure 1. an example of the dtsp-cm instance with straight collecting maneuvers in a cargo airdrop mission. target locations are depicted as small green disks, maneuvers are the red straight line segments, and particular trajectories for a reliable cargo drop-off are in green. the blue curves stand for the dubins maneuvers which connect the determined collecting maneuvers into the closed tour. is required to move straight ahead during the dropping phase itself. an example solution of such an object dropping mission is depicted in fig. 1. in this paper, we propose a novel extension of the ordinary dtsp to address the requirement of specific maneuvers for target accomplishment. the introduced problem is called the dubins traveling salesman problem with collecting maneuvers (dtsp-cm). further, we study a special case of the dtsp-cm where collecting maneuvers are straight line segments (denoted as the dtsp-scm) and we propose modifications of the existing approaches to the dtsp to solve this constrained variant of the dtsp-cm. the paper is organized as follows. an overview of the related work with various dtsp approaches 34 http://dx.doi.org/10.14311/app.2016.6.0034 http://ojs.cvut.cz/ojs/index.php/app vol. 6/2016 dubins traveling salesman problem is provided in section 2. the proposed problem is formally introduced in section 3 and modifications of the existing dtsp approaches are proposed in section 4. evaluation of the proposed modifications of the existing approaches is presented in section 5. finally, conclusion remarks are in section 6. 2. related work various approaches have been proposed to address the dtsp that can be broadly divided into three main classes: 1) decoupled approaches, 2) sampling-based approaches, 3) and evolutionary-based approaches. in this section, a brief overview of representative approaches of each particular class is provided to clarify the context for the introduced dtsp-cm and proposed modifications of the particular approaches to solve the constrained variant of the dtsp-cm with straight line collecting maneuvers. probably the first approach to the dtsp is the alternating algorithm (aa) introduced in [1] which can be categorized as a decoupled approach, since it addresses the dtsp in two consecutive steps. first, a sequence of visits to the targets is determined, e.g., by a solution of the underlying euclidean tsp without considering the vehicle motion constraint. in the second step, headings at the target locations are established by applying straight line segments for even segments followed by closing the tour with dubins maneuvers for odd segments. a similar decoupled strategy has been utilized also in the receding horizon based approach [4] with the horizont of two or three target locations ahead which significantly outperforms the original aa in terms of the solution quality. sampling-based approaches [5, 6] represent the second class of existing solutions for the dtsp. these approaches are based on sampling of possible headings into a finite discrete set and the dtsp is transformed into the generalized asymmetric tsp (gatsp) and further into the atsp by noon-bean transformation [7]. the resulting atsp can be then solved by existing solvers such as concorde [8] or lkh [9]. the third class of the dtsp approaches are evolutionary based algorithms [10]. they encoded solutions into individuals in the population and used similar crossover and mutation operators like for the standard tsp. although evolutionary based approaches can be computationally demanding, their main advantage is the ability to search for a global optima. in [11], authors proposed a novel memetic algorithm to improve performance of the evolutionary based approach by utilizing a local optimization method. recently, the evolutionary multi-objective algorithm nsga-ii [12] has been deployed in [13] to address path planning for data collection in a wireless sensor network that is formulated as the dtsp. the authors consider a bi-objective optimization criterion where the first objective is to minimize the tour length and the second objective is to maximize the time when the vehicle is in the vicinity of the sensor. the problem formulation [13] is similar to the herein proposed dtsp-cm in the considering an influence of the executed maneuver type at the target locations; however, the authors of [13] do not explicitly constraint the type of the maneuver and rather optimize a bi-objective function. on the other hand, the proposed dtsp-cm problem formulation explicitly restricts the collecting maneuver to be one of the specified type with the aim to improve robustness of the solution execution and the mission accomplishment. 3. problem statement the proposed formulation of the dubins traveling salesman problem with collecting maneuvers (dtspcm) is motivated by surveillance missions with fixwing aerial vehicles where the goal is not only to visit the target locations but also to execute a prescribed maneuver for each target. similarly to the dtsp, also in the dtsp-cm, the dubins vehicle [2] is constrained to move only forward at the constant speed v and has the minimum turning radius restricted to ρ. the motion of the vehicle can be described as:  ẋẏ θ̇   = v   cos θsin θ u ρ   , |u| ≤ 1, (1) where u is the control input. the state of the dubins vehicle q is represented as a triplet q = (x,y,θ) and q ∈ se(2), where (x,y) ∈ r2 is the vehicle position in a plane and θ ∈ s1 is the vehicle heading at (x,y). the dtsp-cm stands to determine a closed shortest curvature-constrained tour for which all targets are accomplished by executing a maneuver from a specified set of acceptable collecting maneuvers. each target i is defined by its position pi in the plane and a set of collecting maneuvers ti. the dtsp-cm is formally defined for n target locations p = (p1, . . . ,pn) where pi ∈ r2, corresponding collecting maneuvers t = {t1, . . . ,tn}, and the minimal turning radius of the vehicle ρ. in the dtsp-cm, the dubins vehicle is requested to accomplish all given targets by the particular collecting maneuvers joined in a single closed tour satisfying the vehicle motion constraints (1). therefore, it is necessary to determine both the collecting maneuvers m = {τi, . . . ,τn}, τi ∈ ti, and the sequence of the visits to the targets σ = {σ1, . . . ,σn}. notice, a set of possible collecting maneuvers ti can be infinite, e.g., as an arbitrary selection of the heading angle θ from the set θ ∈ (0, 2π〉 in the ordinary dtsp. the final closed tour consists of the particular collecting maneuvers that are connected to a closed tour by the shortest possible paths respecting the vehicle motion (1), i.e., dubins maneuvers. therefore the total tour 35 petr váňa, jan faigl acta polytechnica ctu proceedings (a) . crossing maneuver (b) . dropping maneuver figure 2. examples of straight collecting maneuvers of the dtsp-scm length l given m and σ can be computed as: l(m, σ) = n−1∑ i=1 d(g(τσi ),f(τσi+1 )) + d(g(τσn ),f(τσ1 )) + n∑ i=1 l(τi), (2) where the function f(τi) represents the starting point of the maneuver τi, and g(τi) is the end point of the maneuver, d(g(τi),f(τj)) stands for the length of the dubins maneuver connecting collecting maneuvers τi and τj, and the symbol l(τi) stands for the length of the τi maneuver. the aim is to minimize the tour length in the dtsp-cm, and therefore, the problem is formulated as the optimization problem formalized in problem 3.1. based on the set of collecting maneuvers, the optimization problem can be both discrete or piecewise continuous. the problem is not always fully continuous, since the length of the dubins maneuver is also only piecewise continuous [14]. problem 3.1 (dtsp-cm). minimizem,σ l(m, σ) subject to ∀i : τi ∈ti 3.1. dtsp-cm with straight collecting maneuvers (dtsp-scm) based on the proposed formulation of the dtsp-cm we further investigate its special variant with collecting maneuvers limited to straight line segments with a prescribed length and its relative position to the particular target location. such limited maneuvers are motivated by real scenarios where the vehicle needs to stay at its course while data from sensors are measured or a cargo is dropped to the particular target location. we call this problem the dtsp with straight collecting maneuvers and denote it as the dtsp-scm in the rest of this paper. having the possible fulfilling maneuver as a straight line segment for which the supporting line is passing the target location, the problem is to find the segment orientation for each target, i.e., we need to determine the vehicle heading at each target location as in the ordinary dtsp. the considered straight collecting maneuver can be specified by two parameters that determine its relative position to the target location: 1) the parameter α represents a distance of the maneuver starting point to the target location; 2) and the parameter β that specifies the end of the collecting maneuver from the target location. for both parameters we can consider positive or negative values according to their relative location to target location defined by the vehicle motion. thus, positive values denote the positions of the maneuver endpoints that are on the approaching side to the target locations (the vehicle is approaching the target), while the negative values denote the position of the endpoints for which the vehicle recedes the target location. the length of the maneuver l(τ) is always positive, and therefore, we define it as l(τ) = |α−β|, l(τ) > 0. (3) we consider two types of straight collecting maneuvers. the first type represents a situation where the target location lies directly on the collecting maneuver. such a maneuver is called the crossing maneuver (α > 0 and β < 0) and it is motivated by robotic scenarios to precisely measure some location and where the vehicle cannot change its heading during the sensing phase. an example of the crossing maneuver is depicted in fig. 2a. the second type of straight collecting maneuvers is a so-called dropping maneuver. in this case, the target location is at the maneuver direction, but not necessary on the maneuver itself, α > 0 and β > 0. this type of maneuver can be used for planning object dropping, since the vehicle is required to focus the target while moving on the straight line segment to ensure a precise airdrop location [15]. after the object dropping, the vehicle can continue in an arbitrary direction and does not need to achieve the target location itself. an example of the dropping maneuver is depicted in fig. 2b. formally, the dtsp-scm is a variant of the optimization problem 3.1. in contrast to the general dtsp-cm, the straight collecting maneuvers have a constant length, since the parameters α and β are fixed for the particular problem instance. therefore, the total length l(m) of all collecting maneuvers can be determined l(m) = n∑ i=1 l(τi) = n∑ i=1 |α−β| (4) and its value is constant that can be computed before solving the particular instance of the dtsp-scm. this property enables to consider only dubins maneuvers length while minimizing the tour length, which significantly simplifies possible methods addressing the dtsp-scm. furthermore, it is possible to modify approaches for the dtsp to address the proposed dtsp-smc due to the constant length of the collecting maneuver. proposed modifications for existing dtsp approaches are presented in the next section. 36 vol. 6/2016 dubins traveling salesman problem 4. solution of the dtsp-scm the main advantage of the proposed dtsp-scm over more general the dtsp-cm is that it is possible to solve the dtsp-scm in a relatively straightforward way by approaches developed for the dtsp with only slight modifications. this is possible because the length of the individual collecting maneuver is constant, and thus the sum of all collecting maneuvers (4) is a constant value and it is not needed to consider individual lengths during the optimization process. hence, we can build solvers for the dtsp-scm on the existing methods for the dtsp. the proposed modifications of the selected representative approaches for each of three classes of the dtsp approaches introduced in section 2 are presented in this section. for decoupled approaches, we can still utilize a solution of the euclidean tsp to determine the sequence of visits to the targets. however, we can expect that such a sequence can be farther from the optimum when the collecting maneuvers are getting longer. a modification of the original aa [1] to utilize collecting maneuvers is straightforward. in solving the ordinary dtsp, vehicle headings are determined for each target location. instead of that, we can determine directions of collecting maneuvers using the direction in odd segments. then, we can connect even segments by the dubins maneuvers to create a closed tour, and thus solve the dtsp-scm. the sampling-based approach [5] is another option how to address the proposed dtsp-scm. instead of sampling the vehicle heading at each target location, we need to sample a set of collecting maneuvers corresponding to each particular target location in the dtsp-scm. in [5], the authors propose to create a distance graph which is an instance of the gatsp. for the proposed dtsp variant with collecting maneuvers, the resulting instance of the gatsp has twice the number of vertices. however, the size of the gatsp instance can be reduced back to the original size by a simple transformation as follows. the length of the utilized collecting maneuver is added to the length of the dubins maneuver which together form a continuous path. similarly the evolutionary-based algorithms [10, 11] are also suitable for a straightforward extension to address the proposed dtsp-scm. here, the chromosome representing the individual needs to be modified to contain the collecting maneuver instead of the vehicle heading only. in addition, the evaluation function needs to be modified to respect the total tour length according to (2). 5. evaluation results several problem instances of the dtsp-scm have been solved to demonstrate usability of the newly proposed variant of the dtsp with straight collecting maneuvers. a small quadrotor helicopter has been simulated in the realistic gazebo simulator [16] with a physical model based on the pixhawk flight controller. figure 3. simulated quadrotor helicopter dropping red ball object. an example of graphical output from the simulator is depicted in fig. 3. the helicopter has been requested to fly at the altitude 20 m with the speed 5 m.s-1 and the minimal turning radius has been set to ρ = 7 m to imitate realistic conditions of a real aircraft. the dropping distance has been set to approximately 9.5 m and the used straight collecting maneuvers have the parameters: α = 12 m and β = 8 m. we have conducted two types of evaluation scenarios. the first scenario is designed to validate a precision of the airdrop maneuvers without and with the explicit consideration of the straight collecting maneuvers in the proposed dtsp-scm. the second scenario is designed to evaluate performance of the proposed modifications of the existing solutions for the dtsp in solving instances of the introduced dtspscm. 5.1. precision of the airdrop first, we found solutions of the ordinary dtsp and the proposed dtsp-scm using the sampling based approach with the lkh solver [17]. the evaluation instance is generated for 10 target locations (n = 10) with the relative density d = 0.3 defined as: d = ρ √ n s , (5) where s is the size of the bounding box where all target locations are positioned. we consider solution of 10 random instances with n = 10 solved by each particular approach (dtsp and dtsp-scm) to obtain statistically significant results. hence, 100 simulated intersection of the dropped object with ground have been recorded. in the used simulated environment, a limited precision of the gps and all the motors were considered except of air resistance. the obtained locations of the dropped object are visualized in fig. 4 where a direction of all dropping maneuvers is rotated to the same orientation to draw all locations within the same figure. it can be easily seen that objects dropped without explicit consideration of the straight collecting maneuvers (i.e., a solution of the ordinary dtsp) are often dropped with a longer distance error from the desired target 37 petr váňa, jan faigl acta polytechnica ctu proceedings ● ● ●● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ●● ● ● ● ● ●● ● ● ●● ● ● ● ● ●● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ● ●● ● −10 −5 0 5 −5 0 5 10 xerror [m] y e rr or [ m ] method ● dtsp dtsp−scm airdrop target figure 4. locations of the dropped object to the target location as a result of the performing dtsp solution and a solution of the newly introduced dtspscm. all measurements are normalized by the position of the airdrop and the target location. 0.0 2.5 5.0 7.5 10.0 12.5 0255075100 percent of dataset [%] e rr or d is ta nc e [m ] solver dtsp dtsp−scm figure 5. precision of the airdrop maneuver obtained by solution of the dtsp and the newly proposed formulation of the dtsp-scm location than for the case when the trajectory is found as a solution of the dtsp-scm. the distance of the real position of the dropped object to its expect position at the target location is depicted in fig. 5. the error for the solutions of the dtsp-scm is almost half of the error for the dtsp. 5.2. performance of dtsp approaches in the dtsp-scm the proposed modifications of the existing approaches for the dtsp to address the introduced dtsp-scm have been evaluated in a series of dtsp-scm instances according to the number of target locations n and their density d determined by (5). six algorithms for the dtsp-scm has been compared. in particular, the proposed modifications of approaches: alternating algorithm (aa) [1], local iterative optimization [18], genetic algorithm [10], memetic algorithm [11], and sampling-based approach [6] with the lkh solver [17] (denoted as the sampling+lkh) and with the optimal concorde solver [8], denoted as the sampling+concorde. due to high computational requirements of the genetic, memetic, and sampling● ● ● ● ● ● ● ● ● ● ●● ● ● ● ● ● ● ● ● ● ● ● ● 0 1 2 3 10 20 50 100 number of target positions r at io o f so lu ti on l en gt h to t he r ef er en ce l en gt h method ● ● ● ● ● ● aa lio genetic memetic sampling+lkh sampling+con. figure 6. average ratios of the solution length to the reference length of the best found solution depending on the number of target locations with the relative density d = 0.3 ● ● ● ● ●● ● ● ● ● ●● ● ● ● ● ● ● 0 1 2 0.2 0.3 0.4 density of target positions r at io o f so lu ti on l en gt h to t he r ef er en ce l en gt h method ● ● ● ● ● ● aa lio genetic memetic sampling+lkh sampling+con. figure 7. average ratios of the solution length to the reference length of the best found solution depending on the density of n = 20 target locations based algorithms, their computational time has been limited to 1 hour. for sampling based approach, a series of solutions is found with iteratively increasing number of samples until the time limit is reached [19]. therefore, all the algorithms provide a solution, albeit more demanding approach (i.e., based on the concorde solver) for less samples than faster heuristic lkh solver, which solves more iterations within the same time limit. the considered number of target locations n has been selected from the set n ∈ {10, 20, 50, 100} and the density d has been selected from the set d ∈ {0.2, 0.3, 0.4}. for each particular scenario defined by the n and d, 20 random instances were generated. for each problem instance, the best found solution (among all solutions found by all approaches) has been selected as a reference solution that allows to aggregate the results into a single solution quality indicator as the ratio of the length of the found tour to the length of the reference solution. average ratios and standard deviations (shown as error bars) are presented in fig. 6 and fig. 7 depending on the number of target locations n and for the particular density of the target locations d, respectively. the results indicate that superior solutions 38 vol. 6/2016 dubins traveling salesman problem are found by the sampling-based approaches. the lkh heuristic approach provides better results than the concorde solver because of limited computational time, and thus solutions with finer discretizations are not computed within the dedicated time limit by the optimal concorde solver. 6. conclusion in this paper, we propose a new variant of the dtsp formulation to address specific requirements of the collecting maneuvers to accomplish visitation of each target location. the proposed variant of the dtsp is called the dubins traveling salesman problem with collecting maneuvers (dtsp-cm). further, we considered collecting maneuvers restricted to straight line segments with fixed length and its specified relative position to the target that is denoted as the dtsp with straight collecting maneuvers (dtsp-scm). for the dtsp-scm, we propose modifications of the existing approaches to the dtsp to address the prescribed constrained collecting maneuvers. the proposed modifications of the algorithms have been evaluated in several scenarios including evaluation of the proposed problem formulation to the precision of the cargo delivery in airdrop maneuvers. the presented results indicate that the proposed formulation improves the reliability of the mission accomplishment. moreover, the presented evaluation results also provide an overview of the performance of particular approaches to the dtsp in the dtsp-scm formulation, where the sampling based methods provide the best found solution among the evaluated approaches. acknowledgements the presented work has been supported by the czech science foundation (gačr) under research project no. 16-24206s. the support of grant no. sgs16/235/ohk3/3t/13 to petr váňa is also gratefully acknowledged. references [1] k. savla, e. frazzoli, f. bullo. on the point-to-point and traveling salesperson problems for dubins’ vehicle. in proceedings of the american control conference, pp. 786–791. ieee, 2005. doi:10.1109/acc.2005.1470055. [2] l. e. dubins. on curves of minimal length with a constraint on average curvature, and with prescribed initial and terminal positions and tangents. american journal of mathematics pp. 497–516, 1957. doi:10.2307/2372560. [3] j. le ny, e. frazzoli, e. feron. the curvature-constrained traveling salesman problem for high point densities. in 46th ieee conference on decision and control, pp. 5985–5990. 2007. doi:10.1109/cdc.2007.4434503. [4] x. ma, d. a. castanon. receding horizon planning for dubins traveling salesman problems. in 45th ieee conference on decision and control, pp. 5453–5458. 2006. doi:10.1109/cdc.2006.376928. [5] j. t. isaacs, d. j. klein, j. p. hespanha. algorithms for the traveling salesman problem with neighborhoods involving a dubins vehicle. in american control conference (acc), pp. 1704–1709. ieee, 2011. doi:10.1109/acc.2011.5991501. [6] k. j. obermeyer, p. oberlin, s. darbha. samplingbased path planning for a visual reconnaissance unmanned air vehicle. journal of guidance, control, and dynamics 35(2):619–631, 2012. doi:10.2514/1.48949. [7] c. e. noon, j. c. bean. a lagrangian based approach for the asymmetric generalized traveling salesman problem. operations research 39(4):623–632, 1991. doi:10.1287/opre.39.4.623. [8] d. applegate, r. bixby, v. chvátal, w. cook. concorde tsp solver, 2003. [cited 31 oct 2016]. [9] k. helsgaun. an effective implementation of the lin-kernighan traveling salesman heuristic. european journal of operational research 126(1):106–130, 2000. doi:10.1016/s0377-2217(99)00284-2. [10] x. yu, j. hung. a genetic algorithm for the dubins traveling salesman problem. in ieee international symposium on industrial electronics, pp. 1256–1261. 2012. doi:10.1109/isie.2012.6237270. [11] x. zhang, j. chen, b. xin, z. peng. a memetic algorithm for path planning of curvature-constrained uavs performing surveillance of multiple ground targets. chinese journal of aeronautics 27(3):622–633, 2014. doi:10.1016/j.cja.2014.04.024. [12] k. deb, a. pratap, s. agarwal, t. meyarivan. a fast and elitist multiobjective genetic algorithm: nsga-ii. ieee transactions on evolutionary computation 6(2):182–197, 2002. doi:10.1109/4235.996017. [13] d. g. macharet, j. w. monteiro, g. r. mateus, m. f. campos. bi-objective data gathering path planning for vehicles with bounded curvature. computers & operations research 2016. doi:10.1016/j.cor.2016.07.004. [14] x. goaoc, h.-s. kim, s. lazard. bounded-curvature shortest paths through a sequence of points using convex optimization. siam journal on computing 42(2):662–684, 2013. doi:10.1137/100816079. [15] c. w. hewgley, o. yakimenko. precision guided airdrop for vertical replenishment of naval vessels. in proceedings of the 20th aiaa aerodynamic decelerator systems technology conference, pp. 4–7. 2009. doi:10.2514/6.2009-2995. [16] gazebo robot simulator. [cited 31 oct 2016]. [17] k. helsgaun. lkh solver 2.0.7. www.akira.ruc.dk/~keld/research/lkh. [cited 31 oct 2016]. [18] p. váňa, j. faigl. on the dubins traveling salesman problem with neighborhoods. international conference on intelligent robots and systems (iros) pp. 4029–4034, 2015. doi:10.1109/iros.2015.7353945. [19] p. váňa, j. faigl. on sampling based methods for the dubins traveling salesman problem with neighborhoods. acta polytechnica ctu proceedings 2(2):57–61, 2015. doi:10.14311/app.2015.1.0057. 39 http://dx.doi.org/10.1109/acc.2005.1470055 http://dx.doi.org/10.2307/2372560 http://dx.doi.org/10.1109/cdc.2007.4434503 http://dx.doi.org/10.1109/cdc.2006.376928 http://dx.doi.org/10.1109/acc.2011.5991501 http://dx.doi.org/10.2514/1.48949 http://dx.doi.org/10.1287/opre.39.4.623 http://dx.doi.org/10.1016/s0377-2217(99)00284-2 http://dx.doi.org/10.1109/isie.2012.6237270 http://dx.doi.org/10.1016/j.cja.2014.04.024 http://dx.doi.org/10.1109/4235.996017 http://dx.doi.org/10.1016/j.cor.2016.07.004 http://dx.doi.org/10.1137/100816079 http://dx.doi.org/10.2514/6.2009-2995 www.akira.ruc.dk/~keld/research/lkh http://dx.doi.org/10.1109/iros.2015.7353945 http://dx.doi.org/10.14311/app.2015.1.0057 acta polytechnica ctu proceedings 6:34–39, 2016 1 introduction 2 related work 3 problem statement 3.1 dtsp-cm with straight collecting maneuvers (dtsp-scm) 4 solution of the dtsp-scm 5 evaluation results 5.1 precision of the airdrop 5.2 performance of dtsp approaches in the dtsp-scm 6 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0044 acta polytechnica ctu proceedings 5:44–46, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app railway infrastructure capacity management for ad hoc trains on the sždc network petr nachtigall∗, martin škárek dopravní fakulta jana pernera, univerzita pardubice, studentská 95, 532 10 pardubice ∗ corresponding author: petr.nachtigall@upce.cz abstract. the paper describes basic principles of ad hoc routes assignment. each carrier has its own interface called information system kadr with sždc for capacity requests management. the paper describes all steps which must be done before train departure keywords: informition system, carrier, ad hoc train, railway infrastructure capacaty. 1. introduction for last couple of years is growing volume of request from railway operators to infrastructure manager (sždc) for so called ad hoc trains [1]. these trains are operated out of the regular construction of the timetable. to make this agenda easier, there was developed an information system called isoř kadr. the effects of the implementation of this system were in much faster and streamline assignment of the ad hoc routes. faster is as well the distribution of the time table to all supporting information systems. among the most important systems which are linked with kadr belong isoř (information system for operative control), compost (train composition) and aport (preparedness and composition of the train). during autumn 2015 kadr stopped being just a planning program, but it was implemented into the basic operation of the traffic. via kadr are constructed deflecting routes by traffic incidents and by traffic situation which requires that the train is going on different than basic route. the sždc has also competency to set the request instead of the carrier and assign the route and the capacity in one step. 2. basic information about kadr information system kadr has two main parts. there is a web part and a desktop part. the access rights for both parts of the system are restricted by roles in the system. there are five basic roles: • carrier – in this role are open up all functions for carriers. this role is available only as a web part, • billing – in this role you can use all functions of the program, • oss – one stop shop – construction of the route, • da – dispatcher – construction of the route, • tnack restrictions – construction of the time table by track restriction (available only in desktop part). each user must be authorized before entering the kadr software. for authorization there is used the logserver. on the logserver are saved all types of roles and for each role are assigned users. after successful authorization the logserver returns back the role (or more roles) for the user. the system automatically starts the interface with functions and data for user’s role [2]. 3. web part of the application the web part of kadr is available anywhere on public internet at the website of the infrastructure manager (http://provoz.szdc.cz/kadr). this part is used mainly by carriers pro submitting of their new requests for assignment of the capacity, for activation and deactivation of trains, for statistics and so on. on the figure 1 is a basic screen of the web part for the carrier after log in. the basic screen has some bookmarks with groups of functions: • requests/djř (day time table) – on this bookmark are all routes of the carrier (including routes from the year time table). each row is one train and carrier can see their basic information and the application status. at the end of each row are icons for work with each request. • activation – this bookmark is for changing of the train status. the carrier can activate the train (train will be really operated), deactivate the train (carrier doesn’t use the capacity that day or period), put together some routes and submit a request for re-activation of the detached train. • train preparation – carrier can set the analysis of the train and preparation of the train for the departure. both information are absolutely necessary for train operation. • restrictions – link for domin (database of infrastructure restrictions) application, where can be find actual track restrictions and other incidents. • carriers – the list of carriers is in register of the infrastructure manager (sždc). this list is in the system called kafr (catalogue of carriers). each carrier can find here basic information about other carriers as well as contacts for responsible people. 44 http://dx.doi.org/10.14311/app.2016.5.0044 http://ojs.cvut.cz/ojs/index.php/app vol. 5/2016 railway infrastructure capacity managment for ad hoc trains figure 1. basic screen of the web part [2]. • numbered lists – there is an overview of all available numberings, such as: . numbered list of engines – information from the information system revoz (register of engines) – basic information about engines and special vehicles. . services – there is a numbered list of services, which are offered by infrastructure manager for ad hoc routes. . operations – numbered list of acts, this can be requested by a carrier. . products – numbered list of products, this can be assigned to a request. • archive – archive of data from previous time tables. • statistics – carrier has to available 4 types of statistics: . charges for capacity assignment – carrier can see financial costs for capacity assignment for any period. . set of capacity requests – list of all capacity requests. requests can be filtered or exported according to various parameters. . set of time tables – list of all assigned requests. requests can be filtered or exported according to various parameters. . day set of time tables – day list of all route of trains. trains can be filtered or exported according to various parameters. probably the most important feature of the web part of kadr is for the carrier submitting of the application. the carrier must fill in step-by-step the following information: • schedule – carrier must choose days for which he submits the request. there are three types of requests. type c – less than 3 working days (the train departs less than 3 days from the request). type b – more than 3 working days (the train departs 4 of more days after the request). this distribution is for easier preferences by capacity assignment. • route of the request – carrier finds his requested route on the map. • basic information about a train – carrier must fill in all mandatory data such as: train engines, weight, length, used type of the brake, category and type of the train. • extending data – here are stated other train parameters, transport of dangerous goods, special load, military transport, services on the train. after completing the entire information carrier is allowed to forward the application to the interfaces of the infrastructure manager. next possibility for submit of the request is via data communication, when carrier has its own software which is compatible with kadr. the data interface is in the format of tsi taf. this requires bidirectional communication between kadr and carrier’s information system. this interface shall be in defined format tsi taf with annexes. nowadays is valid version 5.17 and during this year will be implemented new version 2.1. the web part of the application is for the infrastructure manager very similar to the carrier’s web part. the infrastructure manager doesn’t use this part very often, because most of his usually used functions are in the desktop part. 4. desktop part of the application the desktop part of kadr is available only for infrastructure manager. the installation is executed 45 petr nachtigall, martin škárek acta polytechnica ctu proceedings figure 2. desktop part of the application [2]. through the web part. the work with the desktop part is possible only from the sždc intranet. the infrastructure manager is setting up the time table of ad hoc trains in this part of the application. here can be also authorized the carrier’s request for the ad hoc trains or prepare the aids of the time table like dispatch or tabular time table. the basic screen of the desktop part after authorization is at the figure 2. 5. conclusions the information system kadr becomes during its existence one of the most important planning tool for basic operation of the railway transport. without its existence isn’t nowadays possible ride of any train, because kadr secures activation and deactivation of trains and information system isoř on request of kadr establish the route into all other information systems. the sždc offers to all carriers the web part of kadr for free. each carrier pays only charges and fees for capacity assignment and using of railway infrastructure according to the czech law [3]. references [1] document of infrastructare manager "prohlášení o dráze celostátní u regionální". [2] internal materials of the company oltis group a.s. [3] law nr. 266/1994 sb., o dráhách. 46 acta polytechnica ctu proceedings 5:44–46, 2016 1 introduction 2 basic information about kadr 3 web part of the application 4 desktop part of the application 5 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.10.0065 acta polytechnica ctu proceedings 10:65–68, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app landslide reclamation in šárka valley martin vaníček geosyntetika s.r.o., nikoly tesly 3, prague 6, czech republic correspondence: mvanicek@geosyntetika.cz abstract. housing development in šárka’s valley is situated in former brickworks borrow pit for brick-clay. on its north side the borrow pit cutting is up to 8m high with main road on top. this slope was suffering from progressive landslide mainly due to weathering and surface erosion. the reclamation works proposed and afterwards executed consisted of stabilizing the slope by combination of soil nailing and connecting reinforced slope with surcharge bench at its toe. the final reinforced soil structure nicely merged into the surrounding environment. keywords: landslide reclamation, muralex stone facing system, reinforced soil structure. 1. introduction main task of this project was to design and later to execute the reclamation of landslide below the edge of horoměřická street. the slope is part of housing development project top residence šárecké údolí. this development project is situated in the area of former brickworks. the slope that has been progressively collapsing was up to 8m height with the slope as steep as 1:1. the progressive collapse was initiated by climatic degradation of brick clay and surface erosion on the steep slope. 2. geotechnical conditions in the area of the whole development project the geotechnical investigation has been performed. several boreholes and trial pits have been used to determine the geological structure of the area. as is typical for such projects the soil properties have been suggested from the index tests performed on few disturbed samples taken and from the description of soils encountered during exploratory works. in the area of landslide additional gi was undertaken, however again without determination of shear strength of encountered soils. based on the above information simple geotechnical model was determined. on site bedrock from weathered clayey shales was found, categorised as r5 based on czech code of practise. the depth to weathered bedrock was indicated between 1 and 4 m with respect to the borehole location within the slope. the shales were overlaid by clayey sandy silt (brick clay), based on classification cs sandy clay firm to stiff consistency. near the surface a layer with organic content was present. shear strength properties of soils have been selected with respect to eurocode 7 guidelines as very cautious value from tabled values of soil properties in discontinued czech code of practise čsn 731001. for the construction of new earth structure body, it was in the design assumed to use local soils from basement excavation for houses. also for recompacted soils from excavations the tests for determination of shear strength haven’t been performed, therefore their characteristic properties have been selected as very cautious values from tabled soil properties in čsn 752410, czech code of practise for earth dams. selected characteristic soil properties for reinforced soil structures are presented in table 1 below. based on above mentioned investigation typical cross section through land slide has been created and assumed location of slip surface together with individual soil layers has been indicated in fig. 1. ground water level has been found under the slope relatively shallow, about 1m below ground level. in the borehole 5.5 m deep from top of the slope, ground water hasn’t been found. this indicates that ground water level does not copy the ground surface and is rather naturally gradual. 3. proposed solution in accordance with proposed development project the bottom line of reclamation is defined by edge of unbound cycleway. in order to guarantee required stability of the reclaimed slope, it’s inclination is proposed at 1:1.75 1:2.00 with reinforcement in 1 m spaced layers. due to limited space, the slope has to be ended by retaining wall with max. height of 1.6 m. based on main architect requirement the retaining wall should have gabion facing. at the top edge of the slope, it was not possible to do any large excavations, as there is situated highly trafficked street horoměřická. in order to cut through the potential slip surface, driven soil nails in 1m vertical and 1.5 m horizontal spacing inclined by 60° from horizontal have been proposed. these soil nails from 50 × 6.3 mm steel pipe with 1. 5m length have been located into the steps in existing slope excavations created for better connection with new slope. stability of the new slope is maintained by reinforcing geogrids fortrac® mpt, with vertical spacing of 0.5 m or 1 m depending on the position in the slope. 65 http://dx.doi.org/10.14311/app.2017.10.0065 http://ojs.cvut.cz/ojs/index.php/app martin vaníček acta polytechnica ctu proceedings table 1. characteristic soil properties assumed. figure 1. cross section of collapsed slope with indicated assumed geotechnical layers and slipped surface. the length of geogrids is given between slope face and steps in existing ground, where they are connected to the soil nails by nail heads. the base retaining wall with vertical gabion type facing was designed as muralex® stone system. above the base retaining wall drainage channel was created to drain the surface water from the slope. on eastern side the reclamation is framed by the underpass under the horoměřická street with brick cladding of the steep slope above the underpass portal (see visualization on fig. 2). slope above portal wing wall is again retained by max. 3 m high retaining wall with muralex® stone system. the western side of slope reclamation is framed by newly planned staircase. the level difference between existing slope and reclaimed slope towards the planned staircase is proposed as reinforced steep slope (4:1) with muralex® green system, which provides vegetated surface. figure 2. visualization of reclaimed slope integration into the existing terrain design solution has been developed in accordance to eurocode 7, namely based on experience. typical cross section through reclaimed slope is shown in fig. 3. 66 vol. 10/2017 landslide reclamation in šárka valley figure 3. typical cross section through the reclaimed collapsed slope 4. reclamation execution next to design solution the developer and main contractor of the whole housing development asked us to also take care about the execution of reclamation works. the actual conditions have proven the design assumptions, except the shallower depth of weathered shales close to slope base, where we have some problems with nails driving. this issue was easily solved by shortening of the nails, which does not influence the overall stability. in the upper part the soil nails have been driven well behind the assumed slip surface, along which the slope most probably earlier collapsed. another emerged issue in lack of locally available soil was again easily solved by importing of the suitable soil. this imported soil was tested in lab for shear strength and compaction efficiency by proctor standard test including the optimal moisture content for compaction. reclamation was executed according to plan also due to highly permitting weather. wall from reinforced soil was build using wrap around technique, when individual layers have been compacted against mobile formwork (see fig. 4). at each geogrid layer, anchoring elements have been installed for future connection of muralex® facing elements. at these layers were also installed driven figure 4. formwork for proper compaction of reinforced soil structure with wrap-around facing in the facing area. soil nails and connected to the geogrids strips with help of nail heads (see fig. 5). above the base retaining wall new slope was built with the inclination in between 1:1.75 and 1:2.00 based on space availability. the soil was compacted by plate and ditch roller and compaction quality was tested by light dynamic plate test. once the slope was built and ramp for machinery access removed, attached front facing muralex® system was installed (see fig. 6). in locations with vertical facing, the system was filled with stones 63/125 mm and in locations with 4:1 slope facing, sieved organic soil filling was used. lastly the slope 67 martin vaníček acta polytechnica ctu proceedings figure 5. connection of reinforcing geogrids fortrac® mpt and driven soil nails figure 6. attached front facing system muralex® before 63/125 mm stone filling was covered with topsoil and temporary erosion protection mesh from coconut fibres. afterwards the whole slope surface was covered by grass seeds. already figure 7. overall view on the finished reclaimed slope, where grassing can be seen after couple of days the slope started to be nicely grassed (see fig. 7). 5. summary finally it can be concluded that even at confined conditions it is possible to reclaim collapsed slope by appropriate combination of soil reinforcement techniques soil nailing and soil reinforcement by fortrac® mpt geogrids. the final visual perception nicely implemented the reclaimed slope into high standard housing development. the system of attached front facing muralex® creates aesthetic element, which can be during the lifetime of reinforced soil structure repaired or modified from whatever reasons (e.g. accident, change of aesthetic feeling, . . . ) without compromising the main retaining soil structure. 68 acta polytechnica ctu proceedings 10:65–68, 2017 1 introduction 2 geotechnical conditions 3 proposed solution 4 reclamation execution 5 summary acta polytechnica ctu proceedings doi:10.14311/app.2017.8.0005 acta polytechnica ctu proceedings 8:5–7, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app adhesion of biocompatible tinb coating tomas kolegara, ∗, martin matouseka, monika vilemovab, vladimir starya a faculty of mechanical engineering, department of material engineering, czech technical university in prague, charles square 13, 121 35 prague, czech republic b institute of plasma physics as cr, v.v.i. ∗ corresponding author: tomas.kolegar@fs.cvut.cz abstract. preparation of a coating with a high quality requires good adhesion of the film to the substrate. the paper deals with the adhesion of biocompatible tinb coating with different base materials. several materials such as titanium cp grade 2, titanium alloys ti6al4v and stainless steel aisi 316l were measured. testing samples were made in the shape of small discs. those samples were coated with a tinb layer by using the pvd method (magnetron sputtering). onto the measured layer of tinb an assistant cylinder was stuck using a high strength epoxy adhesive e1100s. the sample with the assistant cylinder was fixed into a special fixture and the whole assembly underwent pull-off testing for adhesion. the main result of this experiment was determining the strength needed to peel the layer and morphology and size of the breakaway. as a result, we will be able to determine the best base material and conditions where the coating will be remain intact with the base material. keywords: pvd coating, adhesion, biocompatible materials, pull-off test, titanium, niobium. 1. introduction today's civilized society brings human kind many simplifications for numerous activities in every day life. due to this comfort more civilization illnesses appear, such as obesity, cancer, atherosclerosis, heart attack and of course musculoskeletal diseases including joints and tendons, but also chronic stress and sports injuries. high life expectancy on other hand causes faster aging of tissue and makes surgery of hip, shoulder and other joints necessary. one solution to help with this this health problem is the utilization of implants. nowadays, commonly used implants are made of lower quality material like stainless steel and these implants are coated with a high quality biocompatible material such as tinb and also other titanium alloys. the most important property of the coating is its adhesion to the base material. the main quality of the parameters for adhesion are the substrate structural properties, purity and surface condition at the time of coating formation, the value of the kinetic energy of falling atoms and ions to the surface, and also the forming a coating on the support surface and the other deposition parameters. according to g. rosenman, the minimum adhesion strength of biocompatible coating to the substrate is 15 mpa [1] and cells have better conditions for adhesion, growth and their establishment in the living body. the goal of this experiment is to verify if the coating is capable of withstanding the strain which can be found in the human body. each human body and conditions inside are very individual and results depend on many factors [2]. 2. materials and procedure for the experiment three base materials were used. titanium cp grade 2, ti6al4v and stainless steel aisi 316l. chemical composition of these materials is shown in table 1, 2 and 3 [3]. experimental samples (discs of diameter 14 mm and 4 mm thickness) were coated with tinb (61 wt % of ti and 39 wt % of nb). the pvd magnetron sputtering method was used [4]. final thickness of the coating was measured by calotest (csm, switzerland) and it was 2.4 ± 0.1 µm. the screw according to din 931 with a 10 mm metric thread was used as the assistant cylinder. these two parts were glued together with high strength epoxy adhesive e1100s with a minimum ultimate tensile strength 70 mpa. the glue was cured in a furnace for 1 hour at 140 ◦c. the pull-off test performed on the available samples was according to en iso 4624:2003 [5, 6]. table 4 presents the measured parameters of the experiment, especially the force when the tested sample was broken. ultimate tensile strength and the average value were measured. 3. results of the experiment (morphology and tensile strength) results of the experiments can be seen in figures 1, 2 and 3. note: the most suitable sample was chosen. black spot indicates detachment of coating. 5 http://dx.doi.org/10.14311/app.2017.8.0005 http://ojs.cvut.cz/ojs/index.php/app t. kolegar, m. matousek, m. vilemova, v. stary acta polytechnica ctu proceedings c n mn p s cr mo ni cu si fe 0.022 0.094 1.75 0.021 0.001 17.18 2.71 14.1 0.06 0.35 rest table 1. chemical composition of aisi 316l [wt %] c n o h fe al v ni mo other ti 0.08 0.03 0.25 0.0125 0.3 rest table 2. chemical composition of titanium cp grade 2 [wt %] figure 1. stainless steel aisi 316l. figure 2. titanium cp grade 2. 4. conclusions • average uts on fe-material was measured (22.48 mpa). . coating stayed on sample. • average uts on ti-material was measured (19.30 mpa). . coating was partially removed. • average uts on ti6al4v-material was measured (34.27 mpa). . large part of coating was removed. figure 3. titanium alloys ti6al4v. ultimate tensile strength of epoxy adhesive e1100s specified by manufacturer is 70 mpa. in the experiment, we achieve 34 mpa in the best result (ti6al4v) while all three testing materials should withstand more than 15 mpa which is the minimum adhesion strength according to g. rosenman [1]. each human body is very individual and we must be sure that implants withstand all the processes in the human body. high values of uts approaching the values set up by manufacturer are more than is required. generally, on a surface with high roughness the adhesion is also higher. further experiments will involve changing parameters of the surface roughness, combination of epoxy adhesion and film adhesion. an ultimate tensile strength of coating at least higher than 50 mpa should be achieved. list of symbols f force [n] ra roughness average [µm] uts ultimate tensile strength [mpa] acknowledgements this work was supported by project lo1207, program npu1, from the ministry of education, youth and sports of the czech republic and by czech science foundation project no. 15-01558s. 6 vol. 8/2017 adhesion of biocompatible tinb coating c n o h fe al v ni mo other ti 0.08 0.05 0.13 0.012 0.25 6.5 4.5 rest table 3. chemical composition of ti6al4v [wt %] sample f ra uts average uts stand. deviation [n] [µm] [mpa] [mpa] [mpa] aisi 316l 1 1540 0.06 ± 0.01 19.62 22.48 2.592 1950 24.84 3 1990 25.35 ti (cp2) 1 1040 13.25 19.30 4.942 1490 18.98 3 1990 25.35 ti6al4v 1 3160 40.25 34.27 4.902 2630 33.50 3 2220 28.28 table 4. table of measured parameters references [1] n. eliaz, o. ritman-hertz, d. aronov, et al. the effect of surface treatments on the adhesion of electrochemically deposited hydroxyapatite coating to titanium and on its interaction with cells and bacteria. journal of materials science: materials in medicine 22(7):1741–1752, 2011. doi:10.1007/s10856-011-4355-y. [2] e. eisenbarth, d. velten, m. müller, et al. biocompatibility of β-stabilizing elements of titanium alloys. biomaterials 25(26):5705 – 5713, 2004. doi:http://doi.org/10.1016/j.biomaterials.2004.01.021. [3] s. guo, j. zhang, x. cheng, x. zhao. a metastable β-type ti-nb binary alloy with low modulus and high strength. journal of alloys and compounds 644:411 – 415, 2015. doi:http://doi.org/10.1016/j.jallcom.2015.05.071. [4] k.-t. k. chiang, r. wei. growth morphology and corrosion resistance of magnetron sputtered cr films. surface and coatings technology 206(7):1660 – 1665, 2011. doi:http://doi.org/10.1016/j.surfcoat.2011.09.020. [5] b. jönsson, s. hogmark. hardness measurements of thin films. thin solid films 114(3):257 – 269, 1984. doi:http://dx.doi.org/10.1016/0040-6090(84)90123-8. [6] x. dai, a. zhou, l. feng, et al. molybdenum thin films with low resistivity and superior adhesion deposited by radio-frequency magnetron sputtering at elevated temperature. thin solid films 567:64 – 71, 2014. doi:http://doi.org/10.1016/j.tsf.2014.07.043. 7 http://dx.doi.org/10.1007/s10856-011-4355-y http://dx.doi.org/http://doi.org/10.1016/j.biomaterials.2004.01.021 http://dx.doi.org/http://doi.org/10.1016/j.jallcom.2015.05.071 http://dx.doi.org/http://doi.org/10.1016/j.surfcoat.2011.09.020 http://dx.doi.org/http://dx.doi.org/10.1016/0040-6090(84)90123-8 http://dx.doi.org/http://doi.org/10.1016/j.tsf.2014.07.043 acta polytechnica ctu proceedings 8:5–7, 2017 1 introduction 2 materials and procedure 3 results of the experiment (morphology and tensile strength) 4 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0072 acta polytechnica ctu proceedings 7:72–75, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app semi–automated assessment of micromechanical properties of the metal foams on the cell-wall level nela krčmářováa, b, ∗, jan šleichrta, tomáš doktora, daniel kytýřa, b, ondřej jiroušeka a czech technical university in prague, faculty of transportation sciences, konviktská 20, prague 1, 155 00, czech republic b the institute of theoretical and applied mechanics as cr, prosecká 76, prague 9, 190 00, czech republic ∗ corresponding author: krcmarova@fd.cvut.cz abstract. metal foams are innovative porous material used for wide range of application such as deformation energy or sound absorption, filter material, or microbiological incubation carrier. to predict mechanical properties of the metal foam is necessary to precisely describe elasto–plastic properties of the foam on cell–wall level. indentation with low load is suitable tool for this purpose. in this paper custom designed instrumented microindentation device was used for measurement of cell-wall characteristics of two different aluminium foams (alporas and alcoras). to demonstrate the possibility of automated statistical estimation of measured characteristics the device had been enhanced by semi-automatic indent positioning and evaluation procedures based on user-defined grid. vickers hardness was measured on two samples made from alporas aluminium foam and one sample from alcoras aluminium foam. average vickers hardness of alporas foam was 24.465 hv1.019 and average vickers hardness of alcoras was 36.585 hv1.019. keywords: metal foam, vickers hardness, cell-wall, indentation under low loads. 1. introduction metal foams are biomimetic porous materials with cellular inner structure that find wide range of applications from deformation energy absorption to noise attenuation, where their very high specific stiffness greatly improves overall effectiveness of constructions [1, 2]. homogenization approach has been proposed as a method for prediction of their mechanical properties on both cell-wall level and at macroscale [3– 6]. however for calculation of macroscopic (effective) mechanical properties by homogenization, mechanical characteristics at the lower level of the foam’s hierarchical microstructure (i.e. cell-wall level) have to be assessed with high precision and reliability. here microindentation is a suitable tool for assessment of required elasto-plastic material properties (for cellwall thicknesses from few hundreds of microns) with the possibility for extension to statistical estimation when automated indents’ positioning and evaluation procedures are introduced. 2. materials and methods 2.1. specimen description and preparation closed-cell aluminium foams with similar compound al 97.0 %, ca 1.5 %, ti 1.5 % (measured using energydispersive x-ray spectroscopy) pore size 2 − 4 mm and wall thickness 100 − 200 µm sales denominated as alporas® (shinko wire co., ltd., japan) and sample 1 sample 2 sample 3 material alporas alcoras alporas width 48 mm 41 mm 44 mm height 21 mm 21 mm 23 mm thickness 10 mm 11 mm 13 mm table 1. dimensions of the specimen alcoras® (alcarbon, germany) were subjected for the testing [7]. inner structure of aluminium foam is depicted in fig. 1. figure was obtained using scanning electron microscopy mira ii (tescan, czech republic). region of interest was chosen according to avoid large pores and structural defect. the structural defects were mainly connected with alcoras sample. from the delivered slabs cuboids with minimal thickness of 12 mm (to ensure sample integrity) were sectioned using water cooled oscillating diamond saw (isomet 1000, buehler gmbh, germany). dimensions of each specimen are listed in tab. 1. low cutting speed 3 mm · min−1 minimised surface damage. samples were embedded into mounting compound (varikleer, buehler gmbh, germany). grinding and polishing procedure employing silicon carbide grinding discs (320, 800, 1200, 4000 grains per square inch) and diamond suspension (1 µm) was performed to remove 1 mm surface layer which could be influenced by sectioning and to obtain plan-parallel faces 72 http://dx.doi.org/10.14311/app.2017.7.0072 http://ojs.cvut.cz/ojs/index.php/app vol. 7/2017 assessment of micromechanical properties of metal foams figure 1. inner structure of aluminium foam obtained by scanning electron microscopy. with minimal surface roughness necessary for proper indentation. 2.2. indentation testing 2.2.1. indentation device description indentation testing was performed using custom designed indentation device developed and constructed on department of mechanics and material fts ctu. indentation device is suitable for hardness measurement by low loads from 10 n up to 100 n. limitation of the device by applying lower loads are: i) inaccuracy of the load recorded by load–cell and ii) insufficient accuracy of the control of the indentation axis. device consists of three independent motorised axes. two of them are designed for precise positioning of the specimen with accuracy 10 µm. third axis is indentation axis equipped by load–cell with positioning accuracy 1.5 µm. device is controlled by load or displacement using gnu/linux software system linuxcnc with custom made graphical user interface. device is equipped with ccd camera (manta g–504b, avt, germany) with a resolution of 2452 × 2056 px attached to a light microscope (navitar imaging, inc., usa) that provided a magnification of up to 24 ×. the acquisition of the projections was controlled by in-house-built opencv based plug-in integrated to control software [8]. indentation device is depicted on fig. 2. this camera is used for the indent place estimation as well as to make a photo of the indent that is than used for hardness measurement. due to high resolution ccd camera, calibrated indentation tip alignment and precise positioning of the specimen, indentation of the selected location of the specimen’s surface is allowed with high accuracy. for verification of the device accuracy of hardness measurement indentation in calibration hardness plate was performed. device overall error estimated by calibration meafigure 2. experimental setup. surement was 1.3 %. maximal measurement error for indentation devices is according to standard set to 3 %. it can be advantageously used for measuring hardness of metal foams because indentation is possible only in places with a sufficient width and depth of the specimen, which means joints of the walls. 2.2.2. experiment procedure description suitable places for indentation on the sample surface were identified using ccd camera and subsequently indentation was automatically performed at these selected locations. after the indentation image data of each indent were captured. the detail of an indent is depicted on fig. 3. indentation load was set to 10 n and this value was reached in 10 s. first indent was used as a testing indent to find indentation speed to respect the condition of reaching maximal force value in 10 s. this indent was removed from data set and hardness value of this indent was not evaluated. all other indents were displacement controlled with given speed identified by the testing indentation. course of the indentation was therefore: i) 10 s loading up to 10 n, ii) 10 s holding on the indentation load and iii) unloading with the same speed as the loading phase. on each specimen was created a series of about 50 indents on the interconnections of the walls. the foam microarchitecture with poresize 2 − 4 mm ensure avoiding that the plastic zones of the indents don’t affect each other. because of size of interconnection and limited minimal indentation force normative 2.5× diameter distance from sample edge can’t be meet during experimental procedure. size of thus generated imprints of the indenter was about 270 µm and their 73 n. krčmářová, j. šleichrt, t. doktor et al. acta polytechnica ctu proceedings figure 3. right: marked area of the imprint obtained by semi–automatic evaluation method. depth was about 39 µm. 2.2.3. evaluation of the experimental data positioning accuracy of the indentation axis does not allow automated evaluation of the hardness from the full position-load curve using oliver–pharr method [9]. therefore hardness was assessed using traditional method by measuring the imprint dimensions and evaluated as vickers hardness: hv = f a ≈ 0.1819f d2 , (1) where f is indentation load in n, a is resulting indentation area of the imprint of the indenter in mm2 and d is average length of the diagonal of the imprint of the indenter in mm. as indentation load is taken the maximal force value registered by the load– cell. length of the diagonals in pixels was measured from the image of each imprint of the indenter using semi–automatic method in matlab software. image of the calibration pattern was used for conversion of the diagonal length in pixels to millimetres. highlighted imprint of the indenter is depicted on fig. 3 3. results elasto-plastic material properties of two types of aluminium foam was measured by vickers hardness. indentation was performed on cell walls and their connections to ensure sufficient place for indentation. in order to get precise information about micromechanical properties of the foam about 50 indents were carried out at each of the three specimens. due to foam nature of specimens the under surface area can be formed by cavity which is not visible on the surface and thus some indents were deformed and these vickers hardness values are omitted from the data set. those indents are usually easily recognised by highly deformed shape of indents. success of the indention was about 60 %. hardness values calculated using equation 1 are listed in tab. 2. sample 1 sample 2 sample 3 material alporas alcoras alporas indent 1 24.630 41.507 18.149 indent 2 33.008 35.490 29.839 indent 3 25.204 36.360 24.817 indent 4 23.027 32.990 23.672 indent 5 21.665 31.514 20.633 indent 6 22.818 34.447 22.967 indent 7 22.211 30.213 23.779 indent 8 24.547 30.254 16.826 indent 9 35.845 49.110 15.233 indent 10 33.856 47.130 13.301 indent 11 27.067 45.138 15.093 indent 12 33.404 28.928 24.319 indent 13 25.699 38.507 33.361 indent 14 21.559 24.547 13.418 indent 15 25.510 41.085 25.044 indent 16 27.996 37.927 16.167 indent 17 32.651 42.484 18.410 indent 18 29.958 30.323 19.584 indent 19 31.472 32.506 24.322 indent 20 21.663 43.721 21.050 indent 21 23.802 34.975 24.192 indent 22 30.054 32.557 31.487 indent 23 34.450 28.754 15.765 indent 24 29.011 45.065 26.474 indent 25 21.114 33.645 15.346 indent 26 40.693 24.999 indent 27 39.462 26.053 indent 28 35.017 indent 29 39.386 indent 30 33.801 avarage 21.641 36.585 27.289 table 2. values of measured vickers hardness by indentation load 10 n as can be seen from the hardness values summarised in tab. 2 vickers hardness of alporas aluminium foam is about 30 % lower than vickers hardness of alcoras aluminium foam. 3.1. comparative measurement to ensure reliability of the indentation process with non-standarded indents size digital microscope vhx74 vol. 7/2017 assessment of micromechanical properties of metal foams figure 4. indent surface inspection and indent profile reconstruction equipped by digital microscope. 5000 (keyence, japan) was employed for indented surface inspection and indent profile reconstruction (see fig. 4). indentation depth ≈ 40 µm was obtained from indent profile using keyence proprietary software. this value corresponds to results of the analysis presented in 2.2.2. the profile is not significantly effected by surrounding embedding resin. 4. conclusions vickers hardness measurement was performed on three samples of aluminium foam on cell-wall level. one sample was aluminium foam alcoras and two samples were made of aluminium foam alporas. about 50 indents were created on each sample on the cell– walls or their interconnections with indentation load 10 n. success of the indentation was about 60 % other indents were deformed due to wrong position on the sample surface what was not possible to avoid prior the indentation. indentation was carried out using custom designed indentation device equipped with load–cell and ccd camera. camera was used for identification of appropriate place for indentation and for acquisition of image data of each imprint. vickers hardness was evaluated from image data using semi–automatic procedure in software matlab. average vickers hardness of the sample 1 (alporas foam) was 21.641 ± 5.496hv1.019, where hv is denotation for vickers hardness and value 1.019 indicates indentation load in kgf. average vickers hardness for sample 3, which was also made from alporas, was 27.289 ± 4.731hv1.019. average vickers hardness of the third sample, which was made from alcoras foam, was 36.585 ± 6.061hv1.019. vickers hardness of alcoras aluminium foam is about 30 % higher than vickers hardness of alporas aluminium foam. information about elasto–plastic properties of the aluminium foams alporas and alcoras on cell–wall level can be used for calculation of macroscopic mechanical properties by homogenization. acknowledgements the research was supported by the czech science foundation (research project no. 15-15480s), by czech technical university in prague (research project no. sgs15/225/ohk2/3t/16) and by institutional support rvo: 68378297. keyence is kindly acknowledged for providing of digital microscope vhx-5000. references [1] t. miyoshi, m. itoh, s. akiyama, a. kitahara. alporas aluminum foam: production process, properties, and applications. advanced engineering materials 2(4):179–183, 2000. doi:10.1002/(sici)15272648(200004)2:4<179::aid-adem179>3.0.co;2-g. [2] m. idris, t. vodenitcharova, m. hoffman. mechanical behaviour and energy absorption of closed-cell aluminium foam panels in uniaxial compression. materials science and engineering: a 517(1–2):37 – 45, 2009. doi:http://dx.doi.org/10.1016/j.msea.2009.03.067. [3] p. koudelka, t. doktor, j. valach, et al. effective elastic moduli of closed-cell aluminium foams homogenization method. 11th imeko tc15 youth symposium on experimental solid mechanics 2012 pp. 244–250, 2012. [4] s. li, d. qian, m. l. and. multiscale mechanics and mechanics of biological materials. [5] v. králík, j. němeček. micromechanical properties of porous material based on metal foam. chemicke listy 105(17):s672–s675, 2011. [6] m. geißendörfer, a. liebscher, c. proppe, et al. stochastic multiscale modeling of metal foams. probabilistic engineering mechanics 37:132–137, 2014. doi:10.1016/j.probengmech.2014.06.006. [7] t. fíla, d. kytýř, p. koudelka, et al. micro-mechanical testing of metal foam cell walls using miniature threepoint bending. key engineering materials 586:120–124, 2014. doi:10.4028/www.scientific.net/kem.586.120. [8] g. bradski. the opencv library. dr dobb’s: the world of software development 2000. http://www.drdobbs.com/open-source/the-opencvlibrary/184404319. [9] w. oliver, g. pharr. an improved technique for determining hardness and elastic modulus using load and displacement sensing indentation experiments. journal of materials research 7(06):1564–1583, 1992. 75 http://dx.doi.org/10.1002/(sici)1527-2648(200004)2:4<179::aid-adem179>3.0.co;2-g http://dx.doi.org/10.1002/(sici)1527-2648(200004)2:4<179::aid-adem179>3.0.co;2-g http://dx.doi.org/http://dx.doi.org/10.1016/j.msea.2009.03.067 http://dx.doi.org/10.1016/j.probengmech.2014.06.006 http://dx.doi.org/10.4028/www.scientific.net/kem.586.120 acta polytechnica ctu proceedings 7:70–73, 2017 1 introduction 2 materials and methods 2.1 specimen description and preparation 2.2 indentation testing 2.2.1 indentation device description 2.2.2 experiment procedure description 2.2.3 evaluation of the experimental data 3 results 3.1 comparative measurement 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0016 acta polytechnica ctu proceedings 13:16–19, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app parallel approach to solve of the direct solution of large sparse systems of linear equations michal bošanský∗, bořek patzák czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: michal.bosansky@fsv.cvut.cz abstract. the paper deals with parallel approach for the numerical solution of large, sparse, non-symmetric systems of linear equations, that can be part of any finite element software. in this contribution, the differences between the sequential and parallel solution are highlighted and the approach to efficiently interface with distributed memory version of superlu solver is described. keywords: linear equation solver, domain, parallelization, threads, distributed memory, communication. 1. introduction many engineering problems lead to large-scale problems. the solution of these problems is being limited by available resources. in many cases, the parallel and distributed computing paradigms can be successfully used to enable solution of large and complex problems, which are not possible to solve on single machine with one processing unit [1]. the traditional serial computers permit to run simulation codes only sequentially and often have limited memory. parallelization can significantly reduce computational time by more efficient use of modern hardware. any parallel algorithm is based on idea of the partitioning the work into the set of smaller task which can be solved concurrently by the simultaneous use of multiple computing resources. parallel computers can be classified, for example, by type of memory architecture. the shared, distributed, and hybrid memory systems [2] exist. in the shared memory system, the main memory with a global address space is shared between all processing units which can directly address and access the same logical memory. the global memory significantly facilitates the design of parallel program, but the memory bus performance is limiting factor for scalability with increasing number of processing units. in distributed memory systems, the memory is physically distributed to individual processing units and there is no global address space. when a processor needs to access data on another processor, it is usually the task of the programmer to explicitly define how and when data is communicated. the cost of communication compared to local memory access can be very high, on the other hand, the advantage is that the overall memory is scalable with increasing number of processors. hybrid systems combine the features of shared and distributed memory system, providing global, shared memory for reasonably small number of processing units that are combined into the distributed memory system. the design of parallel algorithms requires the partitioning of the problem into a set of tasks, that can be processed concurrently. the partitioning of the problem can be either fixed (static load balancing) or can change during the solution (dynamic load balancing) [3]. the latter option is often necessary to achieve a good load balancing of individual processors and consequently also reasonable scalability. the scalability of computation is often regarded as one of the most important goals of parallel computing. the parallel algorithm is considered as scalable, when by using increasing number of processing threads, the execution time is monotonically decreasing, ideally in a linear trend. however, when individual computing threads are mutually dependent, the ideal scalability is difficult to obtain due to the overhead of parallel algorithm (necessary synchronization and communication between threads that is not present in serial code) and due to the fact that some parts of the overall algorithm are essentially serial. in addition to the achieved speedup, the parallel computing allows to solve large, complex problems that do not fit into a single, even well equipped machine. the superlu library provides parallel implementation of the direct solver [4]. it provides serial, multithreaded (shared-memory), and distributed memory version. the implementation and performance of the multithreaded version has been reported by the authors in [5, 6]. in this papers, the focus is on distributed version of superlu based on message passing interface (mpi) [7]. numerical modeling using high performance computers brings in new opportunities. many engineering problems lead to solution of sparse linear system of equations. this is also the case of the finite element method (fem) which is used here as an representative example. in fem the differential equations are converted to the algebraic system of equations by using variational methods with the help of decomposition of the problem domain into sub-domains called elements 16 http://dx.doi.org/10.14311/app.2017.13.0016 http://ojs.cvut.cz/ojs/index.php/app vol. 13/2017 direct solution of large sparse systems of linear equations 68 0 2 11 00 0 3 5 09 2 6 0 04 0 0 13 17 0 0 3 k = k_nzval 8 6 2 11 3 5 9 2 6 4 13 7 1 3 k_colind 1 2 4 5 4 5 1 3 4 1 5 1 2 3 k_row_ptr 1 5 7 10 12 21 3 4 5 1 2 3 4 5 figure 1. the compressed row format of matrix schema. and smart choice of interpolation functions. for linear elasticity problem, considered here, the resulting algebraic equations, represent the discrete equilibrium equations at nodes of the computational mesh k ∗ r = f, (1) where k is the global stiffness matrix, f is global vector of external forces, and r is vector of unknown displacements. one of the key feature of the fem is that the stiffness matrix is typically positive definite, symmetric matrix with sparse structure, which is a consequence of using interpolation and test functions with limited nonzero support. the efficient algorithm for solving linear system must exploit the symmetry and sparse structure, by saving considerable memory and cpu resources. there exist number of different storage schemes for sparse matrices. the most widely used are skyline, compressed row, and compress column formats. in this contribution, the compressed row format will be used. this choice comes from superlu library, which requires the sparse matrix in this format on input. in the compressed row format, only nonzero entries of the sparse matrix are stored in one dimensional array. additional integer array is needed to store column indices of the stored values. in practical implementation, the one dimensional arrays storing nonzero values are merged together into a single array containing all nonzero entries, row by row. the same applies for arrays of column indices. an additional array with size equal to number of rows is needed to point to the beginning of each row record, see fig. 1. 2. implementation superlu interface as noted already in the previous section, the distributed memory programming model does not provide global address space and global memory accessible by all processing nodes, but the memory is distributed among the processing units. for large problems, it is therefore essential to establish a distributed representation of the sparse matrix. this feature is also supported by superlu library, allowing the user to split the global sparse matrix stored in comprow format into blocks, representing compressed row storage for consequent, non-overlapping groups of rows, which are distributed across computing nodes. we assume, that the decomposition of the discretized problem domain has been established and individual, non-overlapping sub domains (partitions) are assigned to and stored locally on individual computing nodes. the so called node-cut strategy is assumed, where the cut dividing the problem domain runs through the nodes. nodes on mutual partition boundaries are called shared nodes, the nodes inside individual partitions are so-called local nodes. each processing node is responsible for assembling its contribution to the global system matrix by summing up the contributions from individual elements. to minimize the communication, it is natural that each processing node will assemble and maintain in its local memory the block of global stiffness matrix rows corresponding to unknowns on its partition. this is uniquely defined for local nodes, which are exclusively shared by local elements on individual partitions. for shared nodes, which are shared by elements from multiple partitions, the ownership has to be defined. in this work, the partition with lowest rank sharing the shared node is the owner of shared node and thus responsible of maintaining the corresponding row entry of the global stiffness matrix. it is clear that for shared nodes, the contributions to the corresponding row have to be received from partitions sharing particular node. the assembly process requires to establish global, unique numbering of equations across computing nodes in addition to the local numbering on individual partitions. the assembly process of the global, distributed system matrix requires basically two steps: set up of data structures required to store the block of compressed row records on individual partitions and the assembly process itself based on localization of individual element contributions into global equilibrium equations according to their connectivity. 2.1. initialization of data structures for distributed matrix this phase should determine the required size of arrays used to store all nonzero entries of the global matrix. the nonzero contributions of each element can be identified from element code numbers. the code numbers represent simultaneously the numbering of equilibrium equations and numbering of corresponding unknowns. the individual values in the stiffness matrix represent nodal forces caused by corresponding unit displacement. the entry in element stiffness matrix representing the nodal force contributing to k-th equilibrium equation and being multiplied by displacement located in global displacement vector at l-th position, should be added to position (k,l) in the global matrix. by assuming that element contributions are full matrices, one can initialize the nonzero structure of the global stiffness matrix using element 17 michal bošanský, bořek patzák acta polytechnica ctu proceedings 1 (0 ,1 ) 1 2 (2 ,3 ) 5 (8 ,9 ) 3 (4 ,5 ) 4 (6 ,7 ) 6 (1 0, 11 ) 7 (1 2, 13 ) 8 (1 4, 15 ) 0 2 3 4 5 6 7 8 9 10 11 12 13 14 15 1 0 2 3 4 5 6 7 8 9 10 11 12 13 14 15 g # 1 # 2 g g # 3 2 5 3 4 6 7 8 # 1 # 2 # 3 4 6 42 # 2 6 k # 1 k # 2 k # 3 -n od e -s ha re d no de # 3 # 3 # 3 # 2 # 2 4 1 f1 f2 10 2 3 4 5 6 7 1 0 2 3 4 5 6 7 32 8 9 6 7 10 11 3 2 8 9 6 7 10 11 76 10 11 12 13 14 15 7 6 10 11 12 13 14 15 { { p1 p2 p3 figure 2. example of simple problem consisting of three elements distributed into different computing nodes, illustrating the global sparse matrix structure and its distribution. 18 vol. 13/2017 direct solution of large sparse systems of linear equations code numbers. in distributed case, each partition is responsible for initializing and prealocating its block structure of compressed rows records. in case of the compressed row storage, the number of nonzero entries in each row has to be determined together with corresponding nonzero column positions of the individual entries. each processing node is responsible for initialization and assembly of the assigned block of rows. this block is composed from local contributions, as well as from shared node contributions from neighboring partitions. therefore, the overall initialization can be divided into (i) local stage, where the local contributions to nonzero pattern are identified (from the partition elements), and (ii) communication stage, where contributions from neighboring partitions are received for locally maintained rows, corresponding to equilibrium equations for shared nodes. any partition has to therefore keep its locally maintained block of compressed rows and also block compressed rows corresponding to shared nodes not being locally maintained. these two blocks can be stored separately. in the initial stage, the block structure is initialized on every partition using only the contributions from local elements. the row entries, that correspond to the shared nodes which are not locally maintained are then communicated (send) to the corresponding partition maintaining corresponding row. after the sending the data, the contributions from neighboring partitions are received containing the contributions to the locally maintained shared equations. non-blocking communication [8] is used to send/receive the contributions. non-blocking communication allows to utilize the potential overlap of communication and computation, which is utilized in the actual implementation. after finishing the communication stage, each partition has fully initialized data structure for locally maintained block of rows. 2.2. assembly of the global distributed matrix after finishing the initialization step, in which memory has been allocated for every non zero entry of the global distributed stiffness matrix, one can proceed with the actual assembly of the matrix from the element contributions. this is done in a similar fashion to the previous step. on each partition the contributions from local elements are assembled. then, the rows corresponding to shared nodes not being locally maintained have to be sent to the partition maintaining corresponding row. finally, each partition has to receive remote contribution to the locally maintained rows of shared nodes. this process is illustrated on fig. 2. in this example, the problem is composed of three elements (corresponding to sub domains as well) and eight nodes, from which five are local on individual partitions and three nodes are shared. 2.3. assembly of the global right hand side the process of assembling the global right hand side vector is very similar to the process of assembling the stiffness matrix, but in many aspects is simpler, do to the assembly of local vector contributions. the global vector is again distributed among partitions, the distribution corresponds to the distribution of matrix rows. the contributions from local elements are assembled first, then those corresponding to rows maintained on remote partitions are communicated and contributions from neighbors are received and entries updated. by finishing the steps described above, the distributed sparse matrix, represented by locally maintained blocks of row entries on individual partitions, can be directly passed into superlu routines. 3. conclusions in this work the algorithm of assembling the distributed sparse matrix in row-column format is described, in order to develop the interface to the distributed superlu solver. the evaluation of the performance and scalability is the subject of ongoing work. the authors also aim to compare the efficiency of the distributed solver to the efficiency of the multithreaded superlu version. acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs16/038/ohk1/1t/11 – advanced algorithms for numerical modeling in mechanics of structures and materials. references [1] b. patzak. parallel computations in structural mechanics. české vysoké učení technické v praze, 2010. [2] c. hufhes, t. hughes. parallel and distributed programming using c++. pearson education, 2003. [3] b. patzak, d. rypl. parallel adaptive finite element computations with dynamic load balancing. proceedings of the first international conference on parallel, distributed and grid computing for engineering 2009. [4] x. s. li, j. w. demmel, j. r. gilbert, et al. superlu users‘ guide program interface. http://crd.lbl.gov/ xiaoye/superlu/, september 1999. [5] m. bosansky, b. patzak. evaluation of different approaches to solution of the direct solution of large, sparse systems of linear equations. advanced materials research 1144(1144):97–101, 2016. [6] m. bosansky, b. patzak. on parallelization of linear system equation solver in finite element software. engineering mechanics 2017 1(1):206–209, 2017. [7] p. s. pacheco. a user‘s guide to mpi. univ. of san francisco, 1997. [8] t. saif, m. parashar. unferstanding the behavior and performance of non-blocking communications in mpi. department of electrical and computer engineering, rutgers university, 2002. 19 acta polytechnica ctu proceedings 13:16–19, 2017 1 introduction 2 implementation superlu interface 2.1 initialization of data structures for distributed matrix 2.2 assembly of the global distributed matrix 2.3 assembly of the global right hand side 3 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0069 acta polytechnica ctu proceedings 15:69–73, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app modeling of monocrystalline magnesium microbeam bending jiří němečeka, ∗, jan maňákb, jiří němečeka a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b institute of physics of ascr, na slovance 1999/2, 182 21 prague 8, czech republic ∗ corresponding author: jiri.nemecek.1@fsv.cvut.cz (ph.d. candidate) abstract. this paper presents a numerical simulation of a micro-scale experiment on a magnesium alloy. micro cantilever beams were fabricated using focused ion beam technology in a single crystal of mg. the cantilever beams have dimensions in the order of a few micrometers and a pentagonal cross section. nanoindenter was used for cantilever beam bending and load-displacement curve was received. cantilevers with two different crystallographic orientations were chosen for the experiment. three dimensional numerical fe model with elastoplastic behavior respecting crystal anisotropy was used to fit experimental load displacement curves. strengths and deformation energy were evaluated from the models for each cantilever. keywords: magnesium, microbeams, crystal anisotropy, fe model. 1. introduction magnesium crystallizes in a hexagonal close packed structure (hpc) with various deformation mechanisms such as basal plane slipping {0001} <1120>, prismatic slip {1010} <1120> and pyramidal slip {1011} <1123> or twinning {1012} <1011> [1]. the deformations start from a defect or on boundaries between different crystalline grains. in the first scenario, slipping happens when the crystal planes slide along each other and in the second, twinning is a concentrated shift of atomic positions [2]. magnesium alloys fulfill the demand for a lightweight construction material with good recycling potential. nowadays magnesium is commonly used in automotive, train and aircraft industry for its high tensile strength, low strength to weight ratio and excellent corrosion resistance. however, magnesium also has negative properties such as tendency for cracking and brittle failure which is probably given due to intercrystalline failure and local transcrystalline fracture at twin zone [3, 4]. nanoindentation is a powerful tool which uses a small diamond tip pressed into the material to measure mechanical properties such as modulus of elasticity, hardness and creep from load-displacement curves [5]. nanoindentation usually operates at microscale and is a perfect tool for measuring properties of an individual crystal. besides classical nanoindentaion testing there are possibilities to use nanoindenter tip for bending microscale samples. micrometer samples can be fabricated using focused ion beam (fib) technology. fib instrument is very similar to a scanning electron microscope (sem) but instead of electrons it uses gallium ions to hit the material and cause its milling [6]. fib energy is typically between 10 and 50 kv, with beam currents varying between 1 pa and 10 na [7] and is capable of creating samples with varied geometry, mostly pillars or triangular and pentagonal micro beams. micro-beam bending is advantageous for evaluating tensile strength and fracture energy. in this work, micro cantilevers beams were tested in specific crystal directions where twinning occurs. during the testing, load-displacement curves were measured. also, a simple numerical model with good reproducibility of the curves gained on different crystalline orientations has been developed. figure 1. crystal orientation with respect to loading direction. 2. experimental part a sample from pure magnesium was prepared. the sample was cut by a precise diamond saw and finely polished to achieve average roughness in the range of a few nanometers. ebsd (electron backscatter diffraction) was performed on the polished surface to find its crystal orientation. micro beams were fabricated on the identified places by the fei quanta 3d feg dual beam instrument combining sem and fib. all samples were micro cantilever beams with length of about 22 µm and have a pentagonal cross section with a width of about 3 µm and variable height 69 http://dx.doi.org/10.14311/app.2018.15.0069 http://ojs.cvut.cz/ojs/index.php/app j. němeček, j. maňák, j. němeček acta polytechnica ctu proceedings figure 2. sem images of fabricated micro beam with c crystal orientation: side view (top) and front view (bottom). from 3.5 to 5 µm. two types of cantilevers with two prismatic orientations, labelled c and d, were created, see figure 1. sem images of the fabricated cantilevers are shown in figures 2 and 3. circle marks were created on each cantilever approximately 2 µm away from the free end. at this mark, nanoindentation was performed with a berkovich tip to bend the cantilever. standard displacement controlled tests were performed with a constant increment of displacement up to a maximal value of 3000 nm. in total, four cantilevers were bent for each orientation, however, the curves had the same trend. in this paper were picked only the typical curves. typical experimental load-displacement curves measured are shown in figure 5. in both cases firstly, elastic deformation characterized by reversible response of the lattice to the displacement. secondly, plastic deformation of the crystal, caused mostly by twinning for the cases figure 3. sem images of fabricated micro beam with d crystal orientation: side view (top) and front view (bottom). of magnesium crystal orientations, takes place. after reaching a critical strain, twinning causes a release of energy that is demonstrated as a radical drop of load in figure 5. 3. numerical modelling a numerical model was created in ansys workbench software 17.2. a three dimensional model was proposed due to the irregular cross section of the cantilever micro beam and due to previous experience of authors with micro beam modelling [10]. material constants for elastic anisotropy of pure mg were taken from the literature [8, 9] and are given in table 1. for the hexagonal crystal, five independent elastic constants are needed and a sixth is calculated from 70 vol. 15/2018 modeling of monocrystalline magnesium microbeam bending figure 4. directional deformation (z axis) at maximal prescribed displacement 3 µm of the cantilever labelled as c with shown mesh. c11 [gpa] c12 [gpa] c13 [gpa] c33 [gpa] c44 [gpa] c66 [gpa] basal 58.58 25.02 20.79 61.11 16.58 16.78 prismatic 61.11 20.79 25.02 58.58 16.58 20.16 table 1. elastic constant of pure mg for basal and prismatic crystal orientation [8, 9]. 0 1000 2000 3000 0 100 200 300 400 500 displacement [nm] l oa d [µ n ] c experimental d experimental figure 5. typical experimental load-displacement curves of two different crystal orientations measured during cantilever bending. plastic strain [µm µm−1] stress [mpa] c d 0 392 260 0.07 470 410 0.010 495 420 0.015 530 425 0.020 545 430 table 2. plasticity data for the two prismatic orientations of crystal c and d. the equation c66 = 1 2 (c11 − c12). (1) the constants are given for basal crystal orientation only and need to be transformed for prismatic orientation. plasticity of the beams was modelled with multilinear isotropic hardening. the yield stress was directly calculated from the bending experiment, where the values for cantilevers c and d are 392 mpa and 260 mpa respectively. the data of plastic strain and stress are shown in table 2. the cantilever is filled with 3-d solid186 elements which are defined by 20 nodes and have three degrees of freedom per node, translation in the nodal x, y, and z directions. this element supports von mises plasticity and is suitable for large deflection [11]. the boundary conditions were set similar to the real cantilever with fixed support on one end and a displacement increment of up to 3000 nm. the deformed shape of the cantilever model is shown in figure 4. the deformation process of cantilever bending can be characterized with deformation energy. a part of the total deformation energy that occurs before the first twinning was calculated from the stress-strain curves as follows g = ∫ �crit 0 σd�, (2) where σ is the axial stress and �crit is the critical strain when the first twinning happens. 71 j. němeček, j. maňák, j. němeček acta polytechnica ctu proceedings 0 1000 2000 3000 0 100 200 300 400 500 twinning displacement [nm] l oa d [µ n ] c experimental c fit figure 6. experimental and calculated loaddisplacement curves of c prismatic crystal orientation with marked the first twinning. 4. results and discussion calculated load-displacement curves can be seen in figures 6 and 7. in both cases fitting curves exhibit the same behavior as in the experiment on the elastic part that matches perfectly. the curves are also with good agreement after yield point until the first twining which is observed on the experimental curves when a radical drop of load occurs. unfortunately this drop brings very difficult task for modelling and it was not considered in this work. according to research on austenitic high managanese steels showing twinning-included plasticity by c.y. sun et al. [12], it would be necessary to employ crystal plasticity incorporating the twinning mechanism, where eighteen material constants are needed to be obtained experimentally. the first twinning, labelled as wcrit, occurs at displacement of 1.2 µm for c and 2 µm for d orientation. the shape of the fitting curves after twinning has the same trend and would match well with the experimental curve if no twinning had happened with only one exception. the slope of the un-loading curve for d orientation is slightly different. the axial stress in the x axis direction is shown in figure 9. the strain was calculated as the equivalent total strain composed from elastic strain and equivalent plastic stain computed of von mises criterion. stress-strain curves calculated at the top fixed end of the cantilever are shown in figure 8. however, this diagram is again partly correct till the wcrit value; the data after this point are not relevant. thus, critical strain (�crit) and stress were deducted from the model. the normal stress till the first drop of load is 619.9 mpa with a strain of 0.0323 for the c cantilever and 592.9 mpa with value of strain 0.0548. the stress-strain curves made it possible to compare the deformation behavior of micro-beams with different orientations. based on this comparison, the 0 1000 2000 3000 0 100 200 300 400 500 twinning displacement [nm] l oa d [µ n ] d experimental d fit figure 7. experimental and calculated loaddisplacement curves of d prismatic crystal orientation with marked the first twinning. 0.000 0.032 0.055 0.080 0.120 0 100 200 300 400 500 592 620 700 800 900 strain [-] st re ss [m p a] c d figure 8. calculated stress-strain curves of c and d prismatic crystal orientations with labelled critical stresses and strains. yield point is higher for c orientation while the stored energy before the first twinning is higher for d. the energies calculated according to equation 2 for c and d are 16.09 jm−3 and 28.08 jm−3 , respectively. that makes d orientation more ductile. 72 vol. 15/2018 modeling of monocrystalline magnesium microbeam bending figure 9. deformed shape and axial stress (x axis) at critical displacement of 2 µm with uninformed wireframe of d cantilever beam. 5. conclusions this paper shows numerical modelling of micro cantilever beams made from pure magnesium with two prismatic orientations of the crystals. the elastic part, modelled with five independent anisotropic constants, gave a perfect match to the experiments. inelastic deformations were captured with isotropic plasticity with multilinear hardening law until the first twinning. the twinning mechanism was not captured and left for future work. it was found that the fitting curves have the same trend as experimental curves if the twinning was removed, which suggests that the model is correct and only the twinning mechanism needs to be incorporated using e.g.crystal plasticity by sun et al. [12]. critical displacements wcrit, corresponding to the critical strains and stresses at the time of the first twinning were evaluated. it was found that the critical stress for c orientation is 30 mpa higher compared to d. however, the critical strain is higher for d. thus, the deformation energy stored in the material before the first twinning is higher for d. list of symbols wcrit displacement at first twinning [nm] �crit strain at first twinning [–] σ axial stress (x axis) [mpa] g deformation energy [j m−3] c66 elastic constant fib focused ion beam sem scanning electron microscope ebsd electron backscatter diffraction acknowledgements financial support of the czech science foundation (projects 17-18652s) is gratefully acknowledged. references [1] k. fukuda, y. koyanagi, m. tsushida, et al. activation stress for slip systems of pure magnesium single crystals in pure shear test. materials transactions 58(4):587–591, 2017. [2] n. k. bourne. unexpected twins. physics 9:19, 2016. [3] f. kaiser, k. kainer. magnesium alloys and technology. john wiley & sons, 2003. [4] t. homma, n. kunito, s. kamado. fabrication of extraordinary high-strength magnesium alloy by hot extrusion. scripta materialia 61(6):644–647, 2009. [5] a. c. fischer-cripps. introduction to contact mechanics, vol. 87. springer, 2000. [6] f. a. stevie, l. giannuzzi, b. prenitzer. the focused ion beam instrument. in introduction to focused ion beams, pp. 1–12. springer, 2005. [7] s. reyntjens, r. puers. a review of focused ion beam applications in microsystem technology. journal of micromechanics and microengineering 11(4):287, 2001. [8] r. hearmon. the elastic constants of anisotropic materials. reviews of modern physics 18(3):409, 1946. [9] m. a. kumar, i. j. beyerlein, r. j. mccabe, c. n. tome. grain neighbour effects on twin transmission in hexagonal close-packed materials. nature communications 7:13826, 2016. [10] j. nemecek, j. nemecek, j. manak. fracture properties of cement studied by nanoindentation and fib in press 2018. [11] ansys, inc. ansys reference manual version 17.2, 2017. https://www.ansys.com/. [12] c. sun, n. guo, m. fu, s. wang. modeling of slip, twinning and transformation induced plastic deformation for twip steel based on crystal plasticity. international journal of plasticity 76:186–212, 2016. 73 https://www.ansys.com/ acta polytechnica ctu proceedings 15:69–73, 2018 1 introduction 2 experimental part 3 numerical modelling 4 results and discussion 5 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.21.0001 acta polytechnica ctu proceedings 21:1–4, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app plasma modification of polyvinyl alcohol microfibers to improve cohesion with cement matrix jakub ďurejea, ∗, zdeněk prošek a, b, jan trejbala, pavel tesáreka a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague, czech republic b czech technical university in prague, uceeb, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: jakub.dureje@fsv.cvut.cz abstract. the article describes plasma modifications of the surface of polyvinyl alcohol (pva) microfibers using oxygen and hydrogen plasma in order to improve the properties of the composite material containing modified microfibers, cement and recyclate. five different modification times 30, 60, 120, 240 and 480 seconds were applied. changes on fiber surface were detected by sem analysis, packed cell wettability measurement, and weight loss during modification. the selected durations of plasma treatment were chosen to produce test samples on which the modulus of elasticity was continuously measured and then bending and compression tests were performed. the measured values were compared with the reference samples. oxygen modified fibers behavior is more hydrophilic compare with reference fibers, but hydrogen modified fibres behave more hydrophobic than reference fibers. keywords: plasma fiber modification, fiber surface modification, hydrogen plasma, oxygen plasma, polyvinylalcohol fibers, pva fibers. 1. introduction cement composite materials are widespread in the construction industry due to their advantageous properties and relatively low cost. unfortunately, unreinforced cement composites are typical for their low tensile and bending strength, and brittleness [1]. in the field of cement composites, this disadvantage is reduced by using reinforcement. one of the possibilities of reinforcing cement composites is the use of scattered reinforcing fibers. as early as 3500 years ago, dispersed reinforcements in the form of straw to reinforce unburnt bricks were used. the article deals with synthetic fibers that have been used since the early 1990s. currently, synthetic fibers are among the most commonly used fibers for reinforcing cement composites in practice [2]. the cohesion of fibers with cement matrix in fiber-reinforced cement composites has a significant effect on its properties. fiber-matrix cohesiveness depends on the interfacial transition zone (itz) that transfers the tension to fibers [3]. to increase (or reduce) the coherence of the fibers with the matrix, plasma surface modification can be employed. plasma is referred to as the fourth state of matter. it consists of charged particles positive ions and negative electrons and neutral particles [4]. plasma fiber modifications process standardly use low-temperature plasma because of low material temperature resistance. plasma modifies fiber surfaces mechanically as well as chemically. with the increasing time of plasma modification process, changes on the surfaces are more significant [5],[6]. as reported in many studies, carbon fibers have been successfully modified for application into epoxy binders and polyetherimide (pei) composites [7],[8]. polypropylene fibers have been modified by plasma to apply colors and adhesives to their surface [9]. plasma modification has successfully increased the resistance of sisal fibers in alkaline environments [10]. in this work, polyvinyl alcohol (pva) microfibers were modified to achieve a slip-hardening behavior of the cement composite. therefore, it is necessary to reduce the chemical bond and increase the physical bond between the fiber surface and the matrix. considering that the process of oxygen plasma termination has a time-unstable character, ie., the modified wettability gradually returns to the reference values, but on the other hand the time-stable character of the physical modification, it is actually the ideal type of modification. the fiber has also been modified in hydrogen plasma, which on surfaces of some materials (diamond) induces hydrophobic behavior, which, along with mechanical modifications, could be ideal for these fibers. 2. materials polyvinylalcohol (pva) microfibers made by kuraray were chosen for modification using oxygen and hydrogen plasma. pva fibers, unlike most other types of polymeric fibers, have good wettability, but the physical bond with the cement matrix is insufficient. the mechanical properties of the fibers used are shown in table 1. surface modifications by oxygen plasma were performed using tesla vt 214 device (inductively stimulated plasma), those by hydrogen plasma were per1 http://dx.doi.org/10.14311/app.2019.21.0001 http://ojs.cvut.cz/ojs/index.php/app j. ďureje, z. prošek, j. trejbal, p. tesárek acta polytechnica ctu proceedings fibers company type length[mm] diameter [µm] tensile strength [mpa] modulus of elasticity [gpa] density [g/cm3] pva kuraray rec15 × 8 8 40 1560 40 1,3 table 1. fiber properties. formed in the roth & rau ak 400 (capacitance stimulated plasma). a radio frequency source of 100 w was used for the both instruments. the samples were placed in the chamber of the device. that was subsequently evacuated to pressure less than 22 pa. subsequently, oxygen or hydrogen gas was applied into the chamber and the pressure increased to about 60 pa. then the radiofrequency plasma generator was triggered for 30 s, 60 s, 120 s, 240 s or 480 s. the temperature of the fibers in the chamber gradually increased during the plasma modification, reaching the temperature in the longest time around 50 ◦c. upon completion of the plasma modification, working gas was re-pumped from the chamber and subsequently aerated to atmospheric pressure. based on the results of these measurements, fibers were selected for the test specimens with a modification time of 480 s in oxygen plasma and 480 s in hydrogen plasma. after the modification, the fibers were immediately placed into the sealed bags to prevent air exchange and unpacked just before production of the samples. another fibers with a 480 s plasma oxygen treatment were stored in a laboratory environmental at 22 ± 1 ◦c and relative humidity 50 ± 2 % (table 2). designation fibers modificationtime [s] plasma gas storage pva 480 s pva 480 oxygen sealed bags pva h2 480 s pva 480 hydrogen sealed bags pva 480 s old pva 480 oxygen open bags table 2. fibers modification. a total of 4 sets which each contains 6 samples were made. the dimensions of the test samples were 40×40×160 mm. portland cement cem 42,5r (radotin) and finely ground concrete recyclate were used to produce the test specimens. the composition of all test samples was identical, differing only in the modifications of the fiber surfaces (table 3). designation cement [g] recyclate [g] water [g] pva ref 400 100 205 pva 480 s 400 100 205 pva h2 480 s 400 100 205 pva 480 s old 400 100 205 designation w/(c + r) fibers [%] volume fibers [g] pva ref 0,41 2 7,32 pva 480 s 0,41 2 7,32 pva h2 480 s 0,41 2 7,32 pva 480 s old 0,41 2 7,32 table 3. composition of the samples. 3. experimental methods and results all fibers were weighed by kern with measuring accuracy 0.1 mg before and after their plasma modification. the highest weight loss occurred in the case of fibers modified by oxygen plasma for 480 s. loss was equal to 0.58 % of their original weight. this was caused by the bombardment of working gas. based on the measured data, it can be concluded that the loss of mechanical properties of the fibers (reduction of the fiber cross-section) is almost negligible. the fiber surface was examined using the merlin zeiss electron microscope, which allows the observed object to be magnified up to 2,000,000 times. to observe objects under an electron microscope, the surface of the fiber needs to be conductive. to obtain the conductivity of the surface of the fibers, a thin layer of gold was sputtered on them. sprinkling was carried out using a quorum q150r es with an au thickness of 5 nm. mechanical changes on fiber surfaces relative to reference fibers were observed by electron microscopy at plasma modification time 240 s and 480 s (figure 1). fiber wettability was measured by an indirect method, similar to that called the packed-cell. because of the small fiber diameter, it was not possible to use direct fiber wettability measurement like for macrofibers. the amount of water on fibers after the fibers were dipped was measured. the fibers were placed in a small container with perforated bottom and lid. subsequently, the fibers were compressed using lead weights. the whole of this system was weighed by kern with accuracy 0.1 mg and then immersed up to the water for 30 seconds. then, the container was taken out of the water and weighed after 120 seconds. the container with lead weights was weighed as well. from the measured data, the percentage weight of the water in the container compared to the weight of the fibers was calculated in figure 2. based on the measurement results, fibers modified by oxygen plasma for 480 seconds and those modified by hydrogen plasma for the same time were selected for the production of test specimens. after modification, the fibers were stored in sealed bags or in open bags for 5 days in a laboratory environment at 22 ± 1◦c and a relative humidity of 50 ± 2%. the samples were tested by bending test with a three-point configuration and under pressure at 28 days of age. samples were loaded with a controlled displacement in a heckert model fp100 hydraulic press at a constant rate of 0.4 mm per minute. the remaining parts of the test 2 vol. 21/2019 plasma modification of polyvinyl alcohol microfibers figure 1. comparison of reference (up) and 480 s modified (down) fibre surface by oxygen plasma. figure 2. the weight of water between the fibers. specimens were tested at pressure with loading speed of 0.8 mm per minute. bending strength for samples with oxygen-modified fibers increased by 15 and 36 %, bending strength for samples which contains hydrogen plasma modified fibers decreased by approximately 20% in compare with reference samples (figure 3). compressive strength was reduced by approximately 5% for hydrogen-modified samples. 4. conclusion pva fibers modified by hydrogen plasma for 480 seconds have been shown to have their more hydrophobic behavior due to the chemical effect of this modification. the mechanical effect of this modification, as with the oxygen plasma modification of these fibers, almost does not occur and the loss of bending strength after reaching the maximum is almost the same as for the reference fibers (figure 4). for pva fibers, it may be appropriate to use hydrogen plasma due to its hydrophobic chemical effect, but it is necessary to significantly increase the mechanical effect of the hydrogen plasma modification. increase of the radiofrequency power or extension of the plasma modification time could increase the mechanical effect of modification. figure 3. comparison of flexural strength. figure 4. force-displacement diagram of bending test. acknowledgements this outcome was supported by the ctu in prague under no. sgs16/201/ohk1/3t/11 and by the ministry of education, youth and sports within national sustainability 3 j. ďureje, z. prošek, j. trejbal, p. tesárek acta polytechnica ctu proceedings programme i, project no. lo1605. references [1] z. keršner. křehkost a lomová mechanika cementových kompozitů., 2005. brno. habilitační práce. vut v brně. [2] d. bharathi. the history of fibre reinforcement. [online], 2015. [cit. 2017-10-31], https://www.linkedin.com/ pulse/history-fibre-reinforcement-bharathi-d. [3] a. bentur, s. mindess. fibre reinforced cementitus composites, 2007. 2nd ed., new york: taylor, isbn 978-020-3088-722. [4] rose, margret. definition plasma. [online], 2017. [cit. 2017-10-31], http://whatis.techtarget.com/definition/plasma. [5] s. r. sabren. cold gas plasma surface modification – optimize plastics bonding adhesion. [online], 2012. [cit. 2017-10-31], http://www.adhesionbonding.com/2012/ 04/30/cold-gas-plasma-surface-modification optimize-plastics-bonding-adhesion/. [6] d. chrysostomou. plasma surface modification for the biomedical industry., 2009. presentation. [7] j. moosburger-will, m. bauer, f. schubert, et al. methyltrimethoxysilane plasma polymerization coating of carbon fiber surfaces. surface and coatings technology pp. 223–230, 2017. [cit. 2018-01-06], doi:10.1016/j.surfcoat.2017.01.017. [8] l. eung-seok, l. choong-hyun, c. yoon-soo, et al. effect of hydrogen plasma-mediated surface modification of carbon fibers on the mechanical properties of carbon-fiber-reinforced polyetherimide composites. composites part b: engineering 2016. [cit. 2018-01-06], doi:10.1016/j.compositesb.2016.10.088. [9] r. pulíček. optimalizace procesu hydrofilizace polypropylenového textilu pomocí dbd atmosférické plazmy, 2016. liberec. diplomová práce, technická univerzita v liberci. [10] b. barra, s. santos, p. bergo, et al. residual sisal fibers treated by methane cold plasma discharge for potential application in cement based material. [online], 2015. [cit. 2018-01-06]. 4 https://www.linkedin.com/pulse/history-fibre-reinforcement-bharathi-d https://www.linkedin.com/pulse/history-fibre-reinforcement-bharathi-d http://whatis.techtarget.com/definition/plasma http://www.adhesionbonding.com/2012/04/30/cold-gas-plasma-surface-modification http://www.adhesionbonding.com/2012/04/30/cold-gas-plasma-surface-modification http://dx.doi.org/10.1016/j.surfcoat.2017.01.017 http://dx.doi.org/10.1016/j.compositesb.2016.10.088 acta polytechnica ctu proceedings 21:1–4, 2019 1 introduction 2 materials 3 experimental methods and results 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.18.0015 acta polytechnica ctu proceedings 18:15–19, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app multi-channel control system for in-situ laboratory loading devices václav radaa, b, ∗, tomáš fílaa, b, petr zlámala, b, daniel kytýřa, petr koudelkab a czech academy of sciences, institute of theoretical and applied mechanics, prosecká 76, 190 00 prague 9, czech republic b czech technical university in prague, faculty of transportation sciences, konviktská 20, 110 00 prague 1, czech republic ∗ corresponding author: rada@itam.cas.cz abstract. in recent years, open-source applications have replaced proprietary software in many fields. especially open-source software tools based on linux operating system have wide range of utilization. in terms of cnc solutions, an open-source system linuxcnc can be used. however, the linuxcnc control software and the graphical user interface (gui) could be developed only on top of hardware abstraction layer. nevertheless, the linuxcnc community provided python interface, which allows for controlling cnc machine using python programming language, therefore whole control software can be developed in python. the paper focuses on a development of a multi-process control software mainly for in-house developed loading devices operated at our institute. the software tool is based on the linuxcnc python interface and qt framework, which gives the software an ability to be modular and effectively adapted for various devices. keywords: cnc, controlling, linuxcnc, python interface, python, parallel programming, qt, pyqt, qwt, pythonqwt, labjack. 1. introduction computer numerical control (cnc) is used for automation of a broad range of machines e. g. milling machines, lathes, 3d printers etc. cnc allows for very effective and precise manufacturing of parts with complex shape and has many other advantageous applications in various industrial fields [1]. there are many cnc software solutions, which vary in price, performance or closed-source commercial (siemens, fanuc, labview) and open-source solutions (linuxcnc, arduino primarily for hobby operation). our research group tends to use open-source solutions, because our needs are different from conventional industrial cnc applications [2]. therefore, some properties of the software has to be operationally modified according to our requirements. for this purpose, closed-source commercial cnc software is not suitable for use with our devices. in our applications, several types of actuators are used including stepper, servo-motors and linear voice coil actuator. purpose of the designs can be divided into three groups: • mechanical loading machines (e. g. in-situ loading devices for x-ray computed tomography) • positioning machines (e. g. optics and sample positioning) • sample preparation devices (e. g. automatic grinders) each machine type is equipped with common parts such as actuator, encoder, limit switches and with application specific equipment such as load cell, thermometer etc. the software package introduced in this paper allows for controlling of both the common and application specific equipment and user the interface can be tailored for the specific application owing to linuxcnc [3] python interface [4]. 2. approach in python, parallel programming [5] is a complex task, particularly due to python global interpreter lock (gil). gil prevents multiple threads to access interpreter internals at the same time by serializing the requests. python threads share all types of variables natively which makes them easy to use, but these threads cannot effectively bring any performance increase by utilizing multiple cpu cores. in order to force python interpreter to perform multiple tasks in parallel and utilize multiple cpu cores, separated python processes must be used, which allows for executing tasks of each process simultaneously. to demonstrate python multithreading and multiprocessing performance, a computational test of prime factorization was performed on 10 million integers. the testing scripts [6] was run on a 4-core, 8-thread (intel core i7-4790k @4.4 ghz) machine. the computation time results are shown in figure 1. the results confirm that python multiprocessing threads do not bring any performance increase. in fact, multiprocessing threads decrease the script performance substantially due to overhead caused by thread switches made by 15 http://dx.doi.org/10.14311/app.2018.18.0015 http://ojs.cvut.cz/ojs/index.php/app v. rada, t. fíla, p. zlámal et al. acta polytechnica ctu proceedings figure 1. comparison of multithreading and multiprocessing performance figure 2. inter-process communication scheme python interpreter. on the other hand, multiprocessing processes do increase the performance because of parallelism. however, sharing variables between python processes is in most cases very ineffective. mostly, it is more effective to implement some kind of inter-process communication. python supports multiprocessing library [7] which provides communication between processes using pipes or queues. python multiprocessing pipes are the simplest approach of multiprocessing communication in python. the pipes connect only two processes. one end of the pipe is used to send data by one process, the other end is used to receive the data by the other process. in order to provide connection between more than two processes, multiprocessing queues can be used, but multiprocessing pipes give more performance due to their simplicity. the control software consists of a multi-process core and graphical user interface (gui) process connected by multiprocessing pipes. the core includes four processes, stat poller, data logger, data keeper and command executor. its interconnection is shown in figure 2. the crucial process in the control software core is the stat poller. it is based on linuxcnc.stat() of linuxcnc python interface and extended by methods providing communication with sensors, output log file management etc. stat poller periodically updates machine status variables calling poll() method and obtains values from sensors. obtained data is sent to data logger, data keeper and to gui. data logger receives data from stat poller and saves it periodically to output file. data keeper keeps received data from stat poller in an array and sends the data to gui on request to show the data in a graph. the control software core contains also command executor process which is based on linuxcnc.command() of linuxcnc python interface. it provides execution of dynamically generated python string commands received from gui or stat poller using exec statement. 3. user interface the main objective of the gui is to not affect the control software core performance, so the user interface runs in a process separate from the control software core and is connected with it using multiprocessing pipes. all the pipes leading from control software core are connected to gui core, which handles communication between control software core and gui process itself. the user interface is aimed to be modular and straightforwardly adaptable for various experimental devices. each experimental device uses its specific linuxcnc initializing file, which contains data needed for control software startup. to take into consideration differences of particular devices, the initialization files were extended by specific entries. these entries define active axes of a device, axes equipped with a sensor (measurement axes), axes units etc. on control software startup, initializing file is loaded and user interface is dynamically generated considering specific entries in initializing file and individual experimental device needs. the user interface of the control software is developed using qt framework in cooperation with python binding pyqt [8]. the user interface is designed as a set of plugins (qwidgets), which contributes to straightforward adaptation to new experimental devices and helps to keep the user interface source code well structured. currently, various gui plugins were created to cover specific needs of laboratory devices. emergency stop button, power button etc. are elementary plugins and are common to all devices. other plugins allow for controlling by executing manual data input (mdi) commands based on g-code (rs-274), manual control of positioning axes of the device or sensors management. for proper sensors functionality initializing files of all supported sensors were created. on control software startup, these initializing files are loaded into software internal database. the files consist of numerical values important for appropriate operation of the sensor specified by calibration protocol or user requirement (measurement range, safe overload, amplification, sensitivity). after the system startup user specifies appropriate sensors which are mounted in particular experimental device using plugin for sensors choice. after the sensors are chosen, gui sends a command to control software core to start obtaining data from sensors periodically (e. g. force, temperature etc.). the control software core sends these data to gui and the data is shown in e. g. load cell plugin. all the data is periodically sent to data logger and saved to output 16 vol. 18/2018 multi-channel control system for in-situ laboratory loading devices figure 3. overall view of the control software user interface file. thus a plugin managing data logging was created. it enables the user to export output file either to a static config-given folder or using a save dialog interface. the plugin also enables creating new empty output file leaving the old data in temporary folder to prevent accidental data loss. the control software purpose is to reliably perform mechanical experiments. there was a special plugin created for controlling the displacement-driven and force-driven experiments. the control software provides a plugin allowing for plotting the data during experiment in a live graph. the live graph window is based on qwt (qt widgets for technical applications) library written in c++ and python binding for this library named pythonqwt . the qwt library was chosen for its performance to achieve higher live graph fps. the plugin enables to specify the amount of samples to plot. the refresh rate of the live graph decreases with increasing amount of samples. the optimal refresh rate is calculated automatically based on the number of samples using linear approximation. overall view of the control software user interface is shown in figure 3. 4. applications the modular design of the control software gives it an ability to control various laboratory devices separatedly or combine them in laboratory setups and control multiple devices as a single complex closed-loop controlled system [9]. moreover it is possible to integrate external procedures (e.g. auto-calibration [10]). functionality of the system is hereinafter demonstrated on the setup consisting of uni-axial loading device in arrangement for optical strain measurement where the camera is placed on triaxial table. the experimental procedure is in detail described in [11]. this is achieved by creating linuxcnc custom initializing file for each device and extending it with [rapo] section. the section name comes from the control software name raposoft (rada positioning software). the section contains specific additional parameters. for instance, the additional parameters of a triaxial motorized table are as follows: [rapo] axes\_active␣=␣x␣y␣z axes\_units␣=␣mm␣mm␣mm axes\_measurement␣= lj\_ains\_active␣= lj\_ains\_gain␣= lj\_ains\_invert␣= the motorized table is equipped with 3 active axes (x, y, z), all of the axes units are millimeters and it has no measurement (loading) axis, i. e. all of the axes are used only for positioning. devices with measurement (loading) axes included mostly use various sensors (typically load cells). parameters needed for sensors operation come with "lj" prefix, which refers to the labjack measurement cards (t7-pro, labjack corporation, usa) , used at our institute for establishing connection with the sensors using labjack.ljm library for python. usage of these parameters is demonstrated on uniaxial loading device. 17 v. rada, t. fíla, p. zlámal et al. acta polytechnica ctu proceedings [rapo] axes\_active␣=␣x axes\_units␣=␣um axes\_measurement␣=␣x lj\_ains\_active␣=␣0 lj\_ains\_gain␣=␣51 lj\_ains\_invert␣=␣1 the uniaxial loading device is equipped with a single active axis, the units are micrometers and the active axis is also measurement (loading) axis. the loading axis is equipped with a sensor (load cell) connected to labjack analog input no. 0. the labjack analog input signal can be amplified using ljtick-inamp amplifier and the amplification coefficient of the analog input is given by the gain parameter. the analog input signal can be inverted to set the appropriate load cell polarity while performing tension and compression tests. the motorized table and the uniaxial loading device can be used together as well. the parameters in initializing file look as follows: [rapo] axes\_active␣=␣x␣y␣z␣a axes\_units␣=␣mm␣mm␣mm␣um axes\_measurement␣=␣a lj\_ains\_active␣=␣0 lj\_ains\_gain␣=␣51 lj\_ains\_invert␣=␣1 the laboratory setup consists of 4 active axes in total, three of them are positioning axes of the table with millimeter units, one of them is measurement (loading) axis with micrometer units. parameters for the labjack analog inputs configuration are the same as for the single uniaxial loading device usage. the control software allows for controlling devices equipped with multiple sensors. this can be demonstrated on initializing file of a loading device developed at our institute used for four-point bending tests. the initializing file looks as follows: [rapo] axes\_active␣=␣x␣y axes\_units␣=␣um␣um axes\_measurement␣=␣x␣y lj\_ains\_active␣=␣0␣1 lj\_ains\_gain␣=␣51␣51 lj\_ains\_invert␣=␣1␣1 the device consists of 2 active axes, both of them are loading (measurement) axes with micrometer units. the device is equipped with 2 sensors (load cells) connected to labjack analog inputs no. 0 and no. 1. both of the analog inputs are amplified using coefficient 51 and their signal is inverted. 5. conclusions the presented control software allows for controlling laboratory devices developed at our institute, which are used for mechanical testing in different loading scenarios including time-dependent processes under controlled ambient conditions. it allows to control laboratory devices actuated by stepper, servo-motors, linear voice coil actuator etc. it provides support for various sensors operating on analog signal proportional to the applied excitation voltage (mv/v) e.g. loadcells or thermometers. moreover it is possible to integrate external closed-loop driven procedures. the modular design enables very effective adaptability and extendability which gives the control software capability to be used in a long-term way. acknowledgements the research has been supported by the european regional development fund in frame of the project kompetenzzentrum mechanobiologie (atcz133) in the interreg v-a austria czech republic programme and by operational programme research, development and education in project engineering applications of microworld physics (cz.02.1.01/0.0/0.0/16_019/0000766). references [1] p. smid. cnc programming handbook, third edition. industrial press, inc., 2007. [2] s. anderberg, t. beno, l. pejryd. energy and cost efficiency in cnc machining from a process planning perspective. in s. g. (ed.), sustainable manufacturing, pp. 393–398. springer, 2012. [3] the linuxcnc team. linuxcnc getting started guide. samurai media ltd., 2016. [4] linuxcnc python interface documantation. http://linuxcnc.org/docs/2.6/html/common/ python-interface.html. accessed: 2018-06-15. [5] g. zaccone. python parallel programming cookbook. packt publishing ltd., 2015. [6] e. bendersky. python parallelizing cpu-bound tasks with multiprocessing. https://eli.thegreenplace.net/2012/01/16/ python-parallelizing-cpu-bound-tasks-with-multiprocessing. accessed: 2018-07-03. [7] b. jones, d. beazley. python cookbook recipes for mastering python 3. o’reilly media, 2013. [8] m. summerfield. rapid gui programming with python and qt: the definitive guide to pyqt programming. prentice hall press, 1st edn., 2007. [9] d. kytyr, p. zlamal, p. koudelka, et al. deformation analysis of gellan-gum based bone scaffold using on-the-fly tomography. materials and design 134:400–417, 2017. doi:10.1016/j.matdes.2017.08.036. [10] m. vopalensky, d. vavrik, i. kumpova. optimization of the x-ray tube voltage with respect to the dynamical resolution in radiography and tomography. ndtnet journal 2018-02, 2018. 18 http://linuxcnc.org/docs/2.6/html/common/python-interface.html http://linuxcnc.org/docs/2.6/html/common/python-interface.html https://eli.thegreenplace.net/2012/01/16/python-parallelizing-cpu-bound-tasks-with-multiprocessing https://eli.thegreenplace.net/2012/01/16/python-parallelizing-cpu-bound-tasks-with-multiprocessing http://dx.doi.org/10.1016/j.matdes.2017.08.036 vol. 18/2018 multi-channel control system for in-situ laboratory loading devices [11] t. doktor, i. kumpova, m. sniechowski, et al. influence of printing and loading direction on mechanical response in 3d printed models of human trabecular bone. acta polytechnica ctu proceedings 18(10):24–27 2018. doi:10.14311/app.2018.18.0024. 19 http://dx.doi.org/10.14311/app.2018.18.0024 acta polytechnica ctu proceedings 18:15–19, 2018 1 introduction 2 approach 3 user interface 4 applications 5 conclusions acknowledgements references untitled vol. 5/2016 scientific student conference interoperability of railway transport iricon 2016 may 4, 2016 prague, czech republic organised by: faculty of transportation sciences czech technical university in prague sponsored by: internal grant agency ctu svk 46/16/f6 guest editors: jitka jírová (faculty of transportation sciences, czech technical university in prague) jitka řezníčková (faculty of transportation sciences, czech technical university in prague) local organizing comittee: vít fábera tomáš brandejský jitka jírová dušan kamenický petr koutecký jitka řezníčková jan zelenka scientific comittee: tomáš brandejský (faculty of transportation sciences, czech technical university in prague) pavel hrubeš (faculty of transportation sciences, czech technical university in prague) josef jíra (faculty of transportation sciences, czech technical university in prague) zdeněk kaufmann (tp iži — technology platform, interoperability of railway infrastructure) martin leso (faculty of transportation sciences, czech technical university in prague) jiří masopust (faculty of electrical engineering, university of west bohemia) petr nachtigall (jan perner transport faculty, university of pardubice) otto plášek (faculty of civil engineering, brno university of technology) jindřich sadil (faculty of transportation sciences, czech technical university in prague) richard svoboda (brno university of technology, faculty of civil engineering) jaroslav vašátko (railway research institute, j.s.c., prague) martina vitteková (faculty of transportation sciences, czech technical university in prague) iii acta polytechnica ctu proceedings book title: scientific student conference interoperability of railway transport — iricon 2016. series title: acta polytechnica ctu proceedings. volume: 5. date of issue: october 15, 2016. published by: czech technical university in prague, zikova 4, 166 36 praha 6 – dejvice, czech republic. ič: 68407700. editorial office: ctu — central library, technická 6, 160 80 praha 6, czech republic. acta@cvut.cz . head of the editorial board: zbyněk škvor (ctu in prague, czech republic). editor-in-chief: iva adlerová (ctu in prague, czech republic). editorial board: robin arthur healey (rectorate of ctu in prague, dept. of external affairs, ctu in prague, czech republic), taťána jarošíková (faculty of biomedical engineering, dept. of natural sciences, ctu in prague, czech republic), milan jirásek (faculty of civil engineering, dept. of mechanics, ctu in prague, czech republic), jitka jírová (faculty of transportation sciences, dept. of mechanics and materials, ctu in prague, czech republic), petr jizba (faculty of nuclear sciences and physical engineering, dept. of physics, ctu in prague, czech republic), pavel kalina (faculty of architecture, dept. of theory and history of architecture, ctu in prague, czech republic), zuzana masáková (faculty of nuclear sciences and physical engineering, dept. of mathematics, ctu in prague, czech republic), jan píchal (faculty of electrical engineering, dept. of physics, ctu in prague, czech republic), miroslav sýkora (klokner institute, dept. of structural reliability, ctu in prague, czech republic), radek šulc (faculty of mechanical engineering, dept. of process engineering, ctu in prague, czech republic). graphic design and typesetting: tomáš hejda (faculty of chemical engineering, dept. of mathematics, university of chemistry and technology prague), jan řezníček (faculty of information technology, ctu in prague, czech republic), j. jílek (ctu in prague, czech republic). available on-line at http://ojs.cvut.cz/ojs/index.php/app each article is assigned a digital object identifier doi:10.14311/app.2016.5.〈4-digit article page number 〉 issn 2336-5382 isbn 978-80-01-06022-3 (1st edition) © czech technical university in prague, 2016. reprinting is allowed only with the agreement of the editorial office and with citation of source. iv vol. 5/2016 preface the papers of the scientific student conference interoperability of railway transport – iricon 2016 are concerning new results coming from the project interoperability of railway infrastructure competence network (iricon) which was coordinated by technical university in brno in collaboration with the technological platform "interoperability of railway infrastructure" and was closed in 2015. the project was supported by the operational programme of the ministry of education, youth and sports – education for competitiveness under the priority axis – tertiary education, research and development. the aim of the project was to improve the quality of human resources for research and development and improve training of target groups, consist of students and young professionals in the field of interoperability of railway infrastructure. necessity of the project was supported by the research interests of students and academic staff of technical universities and research institutes. the project enables to involve young experts in advisory groups for members of the governing organs of the european network interoperability, which was established for the preparation and implementation of the interoperability of the european rail system. the project has set up expert groups called the mirror groups, which reflect the activities of the trans-european rail networks in the areas of: (i) infrastructure, (ii) management and safety, (iii) international railway research board, (iv) energy, (v) system solutions and (vi) research. preparing members of the target groups took the form of: (i) traineeships for young researchers at foreign universities or at european institutes aimed in railway interoperability (ii) organizing workshops and training sessions with the participation of domestic and foreign experts, (iii) participation in international conferences on railway transport, (iv) fellowships at leading foreign railway sites, (v) workshops aimed at exchanging the latest information about news in railway transport and (vi) additional training focused on project management in the field of railway research. besides the aforementioned activities there was created portfolio of complementary activities for the target groups aimed at transfer of technology and improving knowledge in the field of control and legislation in the field of research and development and for evaluation of scientific results. the aim of our conference is to exchange knowledge and connections of scientific and technical potential of universities, research and design institutes together with professional potential of construction and manufacturing companies to promote innovation in rail services in accordance with the requirements of technical specifications for interoperability trans-european rail system in subsystems of infrastructure, energy and management and safety. an important aim of the conference is also to contribute to higher education and professional qualifications of future graduates master and doctoral studies, which will enable them to participate on the management, coordination and other activities in the area of interoperability of the european rail system. otto plášek for the scientific comittee v acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0063 acta polytechnica ctu proceedings 11:63–69, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app possible ways of improving railway operation and infrastructure in area between děčín and rumburk pavel purkarta, ∗, david vodáka, tomáš hořeníb, tomáš javoříka a department of transportation systems, ctu in prague faculty of transportation sciences, prague, czech republic b metroprojekt praha a.s., prague, czech republic, graduate of ctu in prague faculty of transportation sciences, prague, czech republic ∗ corresponding author: purkapav@fd.cvut.cz abstract. this paper is dealing with improving the railway transport on axis děčín – rumburk. this journey can be undertaken in two possible ways (as an interstate one or an international one through germany), however the connections between trains are limited (especially problematic is the rumburk junction, where the connection would be desired). the aim of this paper is to improve the connections in the whole area, in order to reach the maximal effectiveness of a railway passenger transport. keywords: rumburk, děčín, ústecký region, railway transport, periodic timetable, node, service conception, travel times reduction. 1. introduction and goals 1.1. researched area the researched area is located in ústecký region near borders with federal republic of germany, particularly in the area between towns děčín and rumburk (mainly north and south from děčín). the most significant towns in the area are děčín, česká kamenice, rumburk, varndorf and šluknov. whole area is also very attractive for tourists – some railway lines are located very close to the borders of national park české švýcarsko. following railway lines can be included into the researched area (numbers according to the timetable 2016/2017 [1], see fig. 1): • railway line 080 in the section česká lípa – jedlová, • railway line 081 děčín – benešov nad ploučnicí – jedlová – rumburk • railway line 083 děčín – bad schandau – dolní poustevna – šluknov – rumburk, • railway line 084 rumburk – panský – mikulášovice with bypass panský – krásná lípa (– rumburk), • railway line 086 in the section benešov nad ploučnicí – česká lípa, • railway line 089 liberec – zittau – rybniště, • railway line rumburk – jiříkov – ebersbach without regular passenger transport. 1.2. reasons behind selecting the researched area main reason for selection of this area was restoration of operation on border crossing dolní pustevna sebnitz in the year 2014. this brought certain changes in service conception in the area. unfortunately, those changes also caused some connection between trains in the station rumburk to disconnect. despite the fact, that some conceptual changes were made after the year 2014, current situation is still not optimal and that is why the problem is still topical. after the opening of border crossing dolní pustevna – sebnitz, most of the trains in rumburk were concentrated to the time period around minute x.30 [1], but it did not created periodic timetable node, because many trains remained without mutual connection due to infrastructural issues. it is also hardly possible to consider it partial periodic timetable node, because connections were actually interrupted, rather than made functional. change in passenger service conception came with the activation of railway station markvartice, where regular crossing of the passenger trains was moved. paradoxically, this made the creation of connections in the railway station rumburk even more matter of distant future. this step creates really close sequence of trains děčín – jedlová – rumburk and děčcín – česká lípa in the section děčín hl.n. – benešov nad ploučnicí that disrupts periodic sequence of trains in this section. therefore, effect of implemented conceptual changes is questionable. if the goal was to disconnect links to ec trains in the station děčín hl. n. (due to their limited reliability – more frequent and higher delays, compared with average values), it is possible to understand those changes. on the other hand, it is also possible to see those actions as a form of resignation to solve problems where they appear. formulation of those reasons made the team of authors to reflect on this issue and they attempted to find a solution. this solution should solve all the problems and also offer attractive pass through the railway 63 http://dx.doi.org/10.14311/app.2017.11.0063 http://ojs.cvut.cz/ojs/index.php/app p. purkart, d. vodák, t. hoření, t. javořík acta polytechnica ctu proceedings figure 1. diagram of railway network in the researched area. junction rumburk for the passengers. importance of this pass through is regional, but still important. 2. understanding and solving the issue solving of the issue was performed in several steps, as it is usual in transport planning: • detailed analysis of the commuting in the researched area in the relation to the railway transport and its possibilities, • specifying the principles of new service conception and important construction points of the periodic timetable • detailed design of service conception, • specifying necessary infrastructural changes to reach service conception 3. solution of the issue 3.1. summary of facts about commuting in the researched area to asses meaningfulness of transit through the railway junction rumburk it is necessary to evaluate regular commuting in the researched area first. to do so, data from the national counting of inhabitants, houses and flats 2011 [2], has been used. the evaluation of data shows following: • from the regional point of view, both commuting to děčín and rumburk are important. the closer to děčín or rumburk a populated area is, the more people are commuting, • part of inhabitants also commutes to smaller cities such as česká kamenice, benešov nad ploučnicí, šluknov and etc. • very significant is commuting to the regional capital ústí nad labem, but commuting to another regional capital – liberec, has also been found in a part of the area (railway line 038), • many inhabitants from the rumburk area are commuting to varnsdorf and we also cannot omit the influence of česká lípa, • it is appropriate to ensure availability of prague for whole area, • breaking branches of railway lines 081, 083 in rumburk does not influence major numbers of passengers (breaking somewhere else can have worse consequences), but their connection may bring significant improvement of transport in the whole area (direction to/from varnsdorf) of transport to larger cities – that makes this idea meaningful. it is crucial to identify transport relations, which would pass through this junction, to defend creation of connections in rumburk. as it has been already 64 vol. 11/2017 possible ways of improving railway operation and infrastructure mentioned – good connections and fast pass through station rumburk may attract many local inhabitants. significant interactions in relations šluknov – varnsdorf, velký šenov – varnsdorf, mikulášovice – varnsdorf, were identified in regular commuting. another relations, which can be considered as long distance commuting, were also identified. improving of pass through rumburk could have positive effect also on these relations. there is a potential of 250 – 300 regular commuters in villages located only near railway lines (with acceptable walking distance from village to railway line), this number is sufficient enough to take this into consideration. however it is necessary to prefer more important transport relations in order to not create a negative impact on them. 3.2. design of the new service conception design of the new service conception was made according to all the periodical timetable rules. the most important passenger line, which should eliminate all problems in the station rumburk, is new passenger line děčín – benešov nad ploučnicí – jedlová – rumburk – dolní poustevna – sebnitz – bad schandau – děčín. it is proposed to connect branches of both lines děčín – rumburk (an interstate one and an international one through germany) and create pass through model in the station rumburk. basic interval is 120 minutes. interval is shortened to 60 minutes in the section děčín – česká kamenice (– jedlová). another section with interval 60 minutes is děčín – schöna. combination of the new line and long-distance trains in the section jedlová – rumburk – šluknov – dolní poustevna also creates interval of 60 minutes (long-distance trains in the section šluknov – dolní poustevna are operated only on work days). it is crucial to establish important line construction points for periodical timetable. key connection in the station děčín is connection to ec trains from/to prague (node in the minute x.00). extra trains should be extended to/from ústí nad labem in the time of afternoon rush hours, as a fast connection to the regional capital (ec trains are operated only in 120 minutes interval). these trains should be operated with minimum stoppings (in the section děčín – ústí nad labem). other important construction point is in the station jedlová in minute x.00 and in the station rumburk in minute x.30 (because of important connections). second line branch from děčín, in direction bad schandau, arrives to děčín in the minute x.00. there are important connections to long-distance trains děčín – ústí nad labem – lovosice – praha and regional trains děčín – most. both railway line branches, which are entering the junction have the connection to/from prague ensured. in one case the new connection is faster than the one in current timetable. there are 4 vehicles necessary to provide base interval of 120 minutes (without reserve). those trains are also capable of ensuring extra trains děčín – schöna. another two vehicles are necessary for extra rush hour trains ústí nad labem – děčín – česká kamenice – jedlová (without reserve). together with designing new circular line, service conception of the lines in the area was also modified. this includes long-distance trains ústí nad labem – liberec and kolín – rumburk – šluknov and regional trains děčín – česká lípa, jedlová – česká lípa or rybniště – varnsdorf – liberec. also new routes of mikulášovice dolní nádraží – panský – rumburk – ebersbach trains were designed, in case of shortening intervals (mikulášovice dolní nádraží – panský – rumburk) or restoring the regular passenger operation (rumburk – ebersbach). and finally routes for a pair of seasonal tourist trains děčín – jedlová – krásná lípa – panský – mikulášovice dolní nádraží has also been found. complex design of whole system is summarized in the network diagram (fig. 2). 3.3. infrastructural changes necessary for the new service conception following infrastructure modifications are necessary for the application of designed timetable: • reconstruction of the railway station rumburk, so it can handle the operation of periodic timetable node at minute x.30, • reducing the travel times at the benešov nad ploučnicí – rumburk branch (reducing time by 2 or 3 minutes, in some parts of the branch reduction could be reached by using modern vehicles) • reducing the travel times at the rumburk – panský – mikulášovice dolní nádraží and rumburk – šluknov – mikulášovice dolní nádraží branches (reducing time by 5 minutes), • reducing the travel times in the section jedlová – česká lípa to the systematical travel time 30 minutes, • reducing the travel times of long distance trains in the section bakov nad jizerou – česká lípa by 10 minutes, • consider necessary modifications of the stations on railway lines 081 and 083, • in case of implementing a pair of seasonal tourist trains, it is necessary to divide section panský – mikulášovice into two blocks. 3.4. reconstruction of the railway station rumburk new timetable brings higher requirements for continuity and stability of train operations in the station rumburk. 65 p. purkart, d. vodák, t. hoření, t. javořík acta polytechnica ctu proceedings figure 2. network diagram summarizing proposal of service conception in the researched area. 66 vol. 11/2017 possible ways of improving railway operation and infrastructure railway station rumburk is junction for four railway lines from four directions with four neighboring railway stations: • krásná lípa (railway line 081), • šluknov (railway line 083), • panský (railway line 084 – currently with regular passenger transport only on weekends), • jiříkov (currently without regular passenger transport – not numbered). designed timetable assumes regular passenger transport on all four railway lines. new timetable is periodical with base intervals 60 or 120 minutes. in addition, full periodical timetable node in the minute x.30 is established in the station rumburk. first 3 railway lines above are connected to the western part of the station. railway line from jiříkov is connected to the eastern part of the station. operation on station head in the western part of the station, where three railway lines are meeting, must be organized very strictly. main reason is implementing periodical timetable, which causes all arrivals and departures to take place in 10 minutes (6 train routes in 10 minutes). due to pass through model many routes are colliding in the area of station head. to ensure the stability of new timetable arrivals of trains from šluknov, krásná lípa a panský must be performed without collision of routes. based on all mentioned conditions reconstruction of the railway station rumburk seems to be the best solution. reconstruction should include at least tracks configuration, platforms and interlocking [3, 4]. fig. 3 shows ideal variant of reconstruction. western station head allows simultaneous arrival from all three directions at the same time. longer platform edges are primary located by tracks, where trains to krásná lípa and šluknov are expected to stop. these trains have potential to be longer than trains on other branches. signal should be placed to the longer edge of platform 3, in order to coupling train sets. platform 1 is extra. this configuration of the railway station is crucial for implementation of new service conception. 3.5. infrastructure modifications for reducing the travel times and possible new stops from infrastructural point of view, this paper offers solution for benešov nad ploučnicí – jedlová – rumburk – šluknov – mikulášovice dolní nádraží section. upgrading this section is crucial for possible new circular line operation. all modifications are mainly undertaken in the sections between railway stations, railway stations are skipped as a kind of "black box" – cutting the travel times is considered, but modifications were not proposed in detail. main infrastructure modifications are changes in geometrical alignment and establishing speed profile for cant deficiency 130 mm. the longest section with new geometrical alignment has length of 9 km. steel sleepers with shape of "y" are used when there is a new geometrical alignment in a vertical curve, in order to avoid extending subgrade. limit values of the parameters of geometrical alignment are usually reached while increasing the maximal speed. maximal/minimal values are reached on exceptionally [4, 5]. also the reconstruction of signalling and control systems is designed in order to establish device with the third safety level. safety level is increased on some level crossing as well. section rumburk – šluknov is not considered in interlocking reconstructions, because current interlocking has already been ungraded to the third safety level. section šluknov – mikulášovice d.n. (– dolní poustevna) is also skipped, because all railway operation is currently organized according to rule sždc d3. table 1 shows reached travel times reduction on sections. travel times has been calculated in software "dynamika 1.7.1.", with railcar 843 as a reference vehicle. some proposed infrastructure modifications are not necessary, because travel times has already been reached in timetable 2016/17, but their introduction may increase the timetable stability. in order to improve railway transport availability some new stops are proposed – some are designed to replace inappropriate ones and some as completely new. for example: village horní habartice is located really close to the railway, so the new stop could be made there. new stop can replace current stop in veselá pod rabštejnem, which is located too far away from inhabited area. establishing new stop rumburkdolní křečany (fig. 4) on the railway line 083 could be also interesting, because this stop would be located in the center of inhabited area. stop rumburk-dolní křečany already exists on the 084 railway line, but second stop could attract more passengers. 4. conclusion during the problem solving process was managed to prove that the attractive connections between trains in the station rumburk can be reached without major infrastructure changes and without disconnecting any other important connections. this allowed formation of service conception not only for děčín and rumburk areas, but also for parts of česká lípa and liberec areas with overlap to germany. the result is the paper, which presents a possible way how to increase the effectiveness of a railway passenger transport in researched area. acknowledgements all the facts presented in the article are based on the results of work on project 12x2zs railway network of the czech republic and europe, which is undertaken on ctu faculty of transportation sciences, department of transport systems. 67 p. purkart, d. vodák, t. hoření, t. javořík acta polytechnica ctu proceedings figure 3. optimal variant of modifications of the railway station rumburk. section length travel time reduction [min]os r section benešov nad ploučnicí – jedlová benešov n. pl. – markvartice 5,4 -0,5 – markvartice – česká kamenice 7,9 -1 – česká kamenice – jedlová 15,4 -2 – section jedlová – rumburk jedlová – rybniště 9,3 -1,5 -1,5 rybniště – krásná lípa 4,8 -0,5 -0,5 krásná lípa – rumburk 6 -0,5 -0,5 section rumburk – mikulášovice dolní nádraží rumburk – šluknov 10 -2,75 -3 šluknov – mikulášovice d.n. 10,6 -1,75 – table 1. travel time reduction on individual sections. figure 4. proposed location of the new stop rumburk-dolní křečany. 68 vol. 11/2017 possible ways of improving railway operation and infrastructure references [1] sždc. aids of theoretical graph of train running. (for researched area between years 2014 – 2016). [2] commuting to work and school according to summation of inhabitants, houses and flats 2011 – ústecký region [online], 2013. [cit. 2015-02-20], https://www.czso.cz/csu/czso/dojizdka-do-zamestnania-skol-podle-scitani-lidu-domu-a-bytu-2011-usteckykraj-2011-b9vyzff32a. [3] čsn 73 4959. platforms and roofs of platforms of state, regional and industrial railways. únmz praha, 2009. [4] čsn 73 6360-1. geometrical characteristics of railway tracks – part 1: layout. únmz praha, 2008. [5] sždc s.o. drawing tables of superstructure. (for researched area). 69 acta polytechnica ctu proceedings 11:63–69, 2017 1 introduction and goals 1.1 researched area 1.2 reasons behind selecting the researched area 2 understanding and solving the issue 3 solution of the issue 3.1 summary of facts about commuting in the researched area 3.2 design of the new service conception 3.3 infrastructural changes necessary for the new service conception 3.4 reconstruction of the railway station rumburk 3.5 infrastructure modifications for reducing the travel times and possible new stops 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0083 acta polytechnica ctu proceedings 12:83–93, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app an intelligent co-driver surveillance system mădălina toma, mirela popa, leon rothkrantza, b, ∗ a intelligent interaction, delft university of technology, mekelweg 4, delft, the netherlands b faculty of transportation sciences, czech technical university in prague, konviktská 20, prague 1, czech republic ∗ corresponding author: l.j.m.rothkrantz@tudelft.nl abstract. in recent years many car manufacturers developed digital co-drivers , which are able to monitor the driving behaviour of a car. sensors in the car measure if a car passes speed limits, leaves its lane, or violates other traffic rules. a new generation of co-drivers is based on sensors in the car which are able to monitor the driver behaviour. driving a car is a sequence of actions. in case a driver doesn’t show one of the actions the co-driver generates a warning signal. experiments in the car simulator torc were performed to extract the actions of a car driver. these actions were used to develop probabilistic models of the driving behaviour. a prototype of a warning system has been developed and tested in the car simulator. the experiments and test results will be reported in this paper. keywords: behaviour analysis, car simulator, surveillance, bayesian reasoning. 1. introduction in recent years we can observe an increase of investments of car producers and researchers in the development of self-driving cars. the first prototypes of such cars have been tested. however, there is a long way to go before such cars appear on the highways and in the cities. it can be expected that at least for the coming 10 years cars with a human driver will prevail. also in the regular driver-car interaction a lot of automatization will take place. navigation systems support the driver to find the shortest route from his starting point to its destination. smart interfaces support the drivers to use their phones, radio and other digital devices in their cars. this paper focuses on the development of systems increasing the safety of the car driver and other road users. at delft university of technology there is a project running on the development of surveillance angels and guardian angels [1], [2]. one of the applications is a digital automated car driver assistant. such an assistant is able to supervise the driving behaviour of a car driver. the system permanently receives the position of the steering wheel, position of the pedals, gear lever, but also information of body movements of the driver, its facial expression and speech and information from the environment. this information is provided by sensors in the car and sensors attached to car parts and (manual-) instruments. the goal of the system is to make assessments of the possible dispositions/situation of the car driver, that is to say a semantic interpretation of the driving behaviour. in case the driver overlooks important information, violates traffic rules or driving rules, the system will generate an alert or support the driver. the next step is that the car takes over the driving of the car, however this is postponed to future work. the full model is based on information of the car driver, but also on the environment and state of the car. from the movements of the car driver we make assessments of the state of the car driver, his goals and intentions. information on the state of the driver can also be assessed by physiological measures and eeg measurements [3],[4]. this study focuses on the assessment of the visual behaviour of the car driver. a car driver can show many body movements. we can order them in specific categories: • core driving movements. these movements are related to driving a car. • peripheral movements. these movements are related to interaction with peripheral devices in the car such as radio, phone, navigation devices. • signal movements. these movements express some message from the driver about his physical or emotional state (stretching the arms, showing a fist) or messages to other drivers on the road. • random movements. these movements have no specific semantic meaning related to car driving such as scratching the nose. core driving movements can be composed of one single action or a sequence of actions. take, for example, the actions of the car driver such as starting a car, parking, speeding up, slowing down, taking a turn. all these actions are characterized by a sequence of information of the car sensors. after an action the next step is to classify the action. it is necessary to realise that the classification is context sensitive. the classification is usual an ambiguous probabilistic process. in case of an ongoing sequence of actions we have to predict the future and compute the probability of possible next steps. it can also happen that a driver 83 http://dx.doi.org/10.14311/app.2017.12.0083 http://ojs.cvut.cz/ojs/index.php/app m. toma, m. popa, l.j.m. rothkrantz acta polytechnica ctu proceedings skips one or more action on purpose or by mistake. take, for example, the case when a driver wants to take over but does not view the mirror to check if the next lane is free. this can result in a dangerous situation and an alert has to be generated. in this paper a rule based system and a bayesian reasoning system will be used to compute the probability of missing or future steps in an action string. by experiments using a driving simulator all possible sequences of actions are computed as well as corresponding probabilities of those actions. in the next section we discuss related work. in section 3 a model of a digital co-driver is presented. in section 4, experiments with the car driver simulator torc are discussed. then we present an analysis of the recorded data in section 5. this paper is concluded with a final discussion. 2. related work according to the statistics published by the national highway traffic safety administration in 2005 a big number of collisions (78%) and near-collisions (65%) are associated with driver inattention. distractions caused by secondary tasks, such as manipulation of the navigation system, radio, or phone, seems to be the main source of the inattention [5]. in [6] merat and jamson present their findings regarding driver’s ability to respond to sudden and unexpected events while driving under normal conditions and also while performing other tasks. their conclusions highlight that the reaction time increases with 200ms when the driver is using in-car systems. head pose estimation reveals important information about the driver’s focus of attention. in [7] doshi and trivedi present their findings regarding head dynamics and eye gaze as important cues in predicting driver’s intent of changing lanes. regarding human pose estimation, currently there are various representation methods of a human in an image, such as silhouettes [8], bounding boxes, or sticks representing the limbs of a person [9][10]. one of the most popular approaches is using the pictorial structures model which stands as a collection of human body parts [11]. building intelligent systems have become a trend in current researches. in driving domain, they have been called advanced driver assistance systems [12]. according to adarsha [13], these systems support the driver’s sensing ability and they are able to track and detect the error/lapses of drivers. assistive intelligent systems are usual a combination of several tools. for example, benoit, [14] developed in his work an assistive driving system for assessing the driver’s fatigue. his system uses two platforms, namely openinterface developed in c++ for signal processing and icare. icare is a conceptual component model for multimodal input/output interaction. thus, the benoit’s system combines multimodal signal processing analysis with a multimodal interaction. our system can be built on a similar system architecture, and it tries to enhance the novice driver’s skills. many other intelligent systems have been developed [15]. the cognitive model of the driver had been investigated in order to implement driver’s mental activity in this system. the driver’s mental activity is analyzed when a driver interacts dynamically with the driving environment. therefore, the driver mental representation of a situation can be illustrated as an instance in the working memory. we also investigate in building a cognitive model of drivers. we analyzed driver’s mental activity as a result of the following activities: gaze and head activities related to the traffic situation, the car controller (steering wheel, pedals, and gear shifter) related to body parts movements, and states of the car on the road. other intelligent systems have been implemented based on a stochastic model. in these cases, driver behavior has been modeled using collected data from the driving simulator. liang used this model in his work [16], to detect driver distraction in real-time. he modeled driver behavior based on parameters extracted from driver’s eye movements, states of the steering wheel and car position on the lane. in a similar way giusti [17] detected driver sleep-attacks by using acquired data from the steering wheel. wang and gao estimated in [18] the state of a car on the road with a rule-based expert system. the expert system typically provides the intelligence of the system, and it fits very well with the requirements of our system. according to the last decade researches, an expert system can be classified in the following six methodologies: rule-based systems, knowledge-based systems, intelligent agent (ia), database methodology, inference engine, system-user interaction. a basic rule-based expert system has the following entities: a knowledge base which contains data from experts, an inference engine which contains the evaluation of decision making components of the system. we estimated the driver’s intent with a rule-based expert system. moreover, our framework developed with a rule-based expert system reasons like a driver providing full assistance for novice drivers in terms of the driving skills enhancement. fletcher proposed an assistive driver system [19], which monitors driver activity using vision sensors. the set of sensors used in fletcher’s system tracks the eye movements and the body parts of the driver. however, it is difficult to estimate driver intentions from the data of the visual sensors. therefore, our system estimates driver’s intention by correlating data from visual sensors with states of cars and information about the driving environment. the gaming industry provides driving simulations which simulate 3d traffic environments. if a simulation of a 3d traffic environment is used, the data about a driving situation can be read easily. in our work, we used torcs simulator because it can be extended and adapted to the proposed system. this simulator is an excellent 84 vol. 12/2017 an intelligent co-driver surveillance system open source tool, and it gave us an opportunity to configure the tracking sensors and the car controller. in prior research, we tried to determine the driver’s interaction intent by analysing only single body parts, e.g. pose estimation [20], gaze detection, or facial expression. compared with these, we take into account more aspects and actions, such as body postures, gaze direction, and head orientation during interaction with the car and the surrounding environment. a lot of intelligent systems that provide assistance to the drivers were presented above. some of them present good solutions. still an intelligent system with a complex reason which trains the novice driver’s skills need to be developed. thus, we present the implementation of an intelligent system that has the capability to assist a novice driver and it tries to enhance his skills at the same time. 3. model in figure 1 we display a model/architecture of our digital co-driver system. the behaviour of the driver is assessed by parallel sensor systems: kinect, gaze tracker, car-driver interaction tools. the sensory input system has been configured and synchronized with clock cycle of torcs tool in order to track the driver activities. the torcs simulator is extended on three modular levels of abstraction. in the first level, the system detects sequences of postures. the driver action can be recognized from sequence of postures in the second level. in the last level, a driver’s intention is estimated based on the driver’s actions related to the traffic situation. the set of possible intentions had been stored in the database. when a driver’s intention is wrong the proposed framework sends feedback alarms. a gui interface has been built in order to allow monitoring the functionality of the whole system and also driver’s activity. the main reasoning module of the system implements the intelligence of the system, and it is responsible to predict driver’s intent. in one of the developed prototypes a rule based system was used to assess the most probable scenario. every rule has a tally attached to it to indicate the probability or importance of the rule. some scenarios are composed of one single action. an observed action triggers the most probable rule. this activated rule is allowed to fire and the right hand site of the if-then rule is generated. some scenarios consist of only one action. if this is a dangerous driving action an alert has been generated by a rule. an example is a driver who doesn’t look at the road in front of him. in other cases observed action(s) trigger the most probable scenario and then specific rules check if important actions are missing or dangerous actions can be expected. an example is a car driver switching on his signal lights to change his lane who did not check of the next lane is free to insert. the problem with rule based systems is that all the rules have to be defined in advance. so the system is not prepared for new, unexpected situations. another problem is that sensors are not able to assess all the ongoing actions. these errors in the sensor observation can be caused by failing technology or because vision field of sensors is occluded. the actions are ambiguous or the lighting conditions are too bad. a gaze tracker, for example, usually misses a significant amount of actions. by sudden, large movements of the head the gaze tracker has to be calibrated again. nevertheless, the biggest problem to solve is that actions are usually distributed in time and that some car drivers start actions simultaneously. it is necessary to realize that the proposed system will be used as a decision support system. ultimately car drivers themselves remain responsible for their actions. to improve the performance of the assessment of the actions a bayesian probabilistic system will be used. to compute the values of the entries in conditional probabilities tables experts are needed to set these values or huge amount of data generated in experiments testing the system is needed to compute the values. the robustness of the system is increased with a fusion method based on multiple sensors compared with a fusion method based on a single sensor. multi-sensor data fusion allows the combination of information with different physical characteristics to enhance the understanding of the driver’s action. the information read from the set of sensors are fundamental in the decision-making module of the system. our fusion method seeks to combine information from multiple sensors. table 1 presents a hierarchy of driver activity in a bottom-up way, and it represents the foundation of our system architecture . the first level needs to detect the driver postures, and then sequences of postures define the driver action. in the third level, one action or a set of actions predict the driver’s intention. a human intention can be defined as an anticipated outcome that guides planned actions. it is the goal or purpose behind an action or a set of actions that a person is following. driver’s intention can be predicted from the driving actions. driving actions in a scenario can be recognized based on a set of postures performed over time. thus, the system architecture depicted in figure 1 defines an assistive car driving system based on multi-steps of processing: tracking driver behavior, predicts driver’s future behavior, decide if the driver’s intention is correct or not, and issue corrective feedback. 4. simulation environment the main tool in our system is torcs simulator. it is one of the most popular 3d open source car simulators written in c++ and available under gpl license. it can be used as an ordinary car racing game, as an ai racing game. the main advantage of the software is that we could extend the platform by configuring inside all sensors used for tracking the driver’s behavior. we selected torcs in our research for the following reasons. torcs is an advanced fully 85 m. toma, m. popa, l.j.m. rothkrantz acta polytechnica ctu proceedings figure 1. architecture of the digital co-driver system. levels of driver activity driver interaction with car controller and traffic situations driving intent interaction: is a mental state of any driver who executes an action or set of actions into a driving situation based on driving environment and the car states. driver’s actions: are made up of multiple body-part gestures such as eyes, head, arms, legs motion, and each body-part gesture is an elementary event of motion and can be composed of a sequence of instantaneous poses at each moment of time. driver’s postures: are instantaneous configuration of the body part in the space of interaction. table 1. hierarchy of driver activity. customizable simulator, and it can be adapted for our application; it features a sophisticated physics engine as well as a 3d graphics engine for the visualization of the virtual environment; and it has a modular software architecture, hence the new controlling and sensing devices allow a straightforward integrating. according to the features of the proposed system, we design tracks which simulate some driving scenarios. therefore, the creation of a customized track involves the use of other tools. in the first step, we used the trackeditor tool for creation a track. it can design the tracks adding straight or curves and some parameters can be configured such as length, radius or banking. all of this information can be stored in an xml file. in addition, the trackeditor tool generates a file with extension. ac, and this file stores the 3d description of the track. the 3d description of the track was edited with the blender tool. blender tool allows us to add elements such as traffic signs and to insert textures. we added billboards to one of the tracks. looking at the driver and understanding his actions is the first task of our system. we used a marker-less system to track upper body activity, and the driver’s upper limb motion can be tracked by the kinect device. this device has been developed by microsoft as a game console input device. comparing with other video cameras [21] used for recognizing human’s movement activities, it tracks the motion of the subject through a combination of hardware and software technologies and achieves a high accuracy tracking of the body with a rate of 30 fps (frames per seconds). in our system, we used for eye movements and head orientation the eyelink ii device produced by the canadian sr research, which uses pupil and cornea reflection tracking mode. this system provides a comprehensive tracking based on smart vision sensors which consist of a head-mounted camera system and two pcs for processing data and running experiments. 86 vol. 12/2017 an intelligent co-driver surveillance system on the head-mounted device, both left and right eye pupil position and the head orientation relative to computer monitor can be tracked. combining the position of the head with the pupil movement relative to the screen, enabled recording of the gaze direction. sensors for physical driver’s interaction. our proposed system (see figure 2, figure 3 and figure 5) presents the architecture of a driver-computer interaction system. the driver can interact naturally with the simulated traffic environments through physical input/output interaction interfaces by bridging the gap between the digital and physical world. figure 2. car driving simulator. figure 3. typical scene from the curved race track with billboards. figure 4. the curved race track driven by the participant. markers indicate locations of the billboards, if enabled. we used as an input interface a logitech g27 joystick controller based on a steering wheel, a gear shifter and three pedals (clutch, brake and throttle). it is similar to the usual basic control system from a cockpit car and the driver’s actions can be performed in the same way as in the real car. the driving scenes are displaying on an output interface which is a large screen like a real car windshield. we used a tv screen with 56 inches size diameter as a large screen. four optical sensors of head camera from the eyelink ii device had placed in the four corners of the screen. 5. experiments in this section we discuss two experiments testing respondents in our driving simulator torc as displayed in figure 5. several devices were used to assess the actions of the driver. in section 5.2 we discuss the assessment procedures using the kinect device. but first we discuss the experimental results of assessment of driving actions using the car sensors as discussed in section 3. data captured with video cameras and gaze tracker has been captured with different sampling rates. the same holds with data from the kinect sensor. recorded data is redundant, full of errors and missing data. the different data streams have to be fused. different types of data fusion using multiple, multimodal streams have been discussed in [22]. figure 5. test person in action in the driving simulator. 5.1. experiment 1 detection behaviour actions in total, there were 23 students invited to take part in the driving experiment. they were supposed to drive for two hours, one hour on a simple trajectory and one hour on a curved trajectory. they had to drive with different speed, normal speed and as fast as possible. along the simple trajectory we placed some billboards. there were other car drivers on the road driving both in a low and high speed speed. this forced our test persons to take over at regular times. we were interested in isolated driving actions or scenarios composed of single driving actions. in figure 6 we show a state diagram of all possible states of actions and transition between states. during driving, the students got phone calls or visual messages about routing. the billboards along the road caught the car driver’s attention. 87 m. toma, m. popa, l.j.m. rothkrantz acta polytechnica ctu proceedings figure 6. the state diagram describing the driver intention and rules used on each transition. 5.1.1. analysis experiment 1 the goal of our system is to generate alerts in case of dangerous driver’s behaviour. we analyse single actions and scenarios composed of actions. a car driver is supposed to watch the road all the time. in case he/she is looking left or right for some time, looking to the billboards of traffic signs for a long time or looking at their telephone or routing devices for a long time, a dangerous situation can occur. assessing the gaze direction via the gaze tracker is rather complicated. fast movements over large view angles were difficult to track. sensors in the car detected if the car left its lane. in the next section we discuss our experiments using kinect to assess the position of the head and gaze direction. parallel to the single action analysis we researched possible scenarios. as soon as the action has been detected we find the position in the state diagram and research if this state is the starting point of a scenario or a point within a scenario. if an important preceding action is missing an alert can be generated. in case of the take-over scenario it can happen that the driver switch on the signal lights but monitoring the side mirror is not detected. to reduce the amount of false alarms we used a probabilistic approach. states in the state diagram and transition between states have probability numbers depending on the frequencies in all recordings during the experiment. we choose the most probable states and transition path to compute the possibility of an alert. we tested 23 drivers during 2 hours (high-low speed). on average, every minute a single action or a scenario of average 3 actions of the driver was required or generated. in total we observed 16.500 single actions and 48.230 single action in scenarios (see table 2). we focused on the actions with most frequent errors. 5.2. experiment using kinect this section presents one of the first digital co-driver experiments [23]. we present adapted behavioral models, which are based on our driving experience and also on information gathered from experts in the field. we assess both head pose and body pose in terms of orientation and hands position in relation to the car environment. we defined normal vs. dangerous driver’s behavior, on several levels of importance. normal driver’s behavior implies both hands on the steering wheel and the head orientation frontal. there are variations from these rules, such as: head orientation can be left or right for a small amount of time in case of lane changing or observation of the surroundings, etc. the dangerous behavior appears in case the driver keeps looking in any other direction than frontal for a long period of time. regarding hands position, the driver is supposed to have at least one hand on the steering wheel, while the other one can rest, be used for changing the gears or 88 vol. 12/2017 an intelligent co-driver surveillance system high speed low speed actions/scenarios frequency errors correct alert frequency correct alert looking around 121 72% 276 81% switching lanes 322 68% 158 77% giving priority 220 75% 118 87% speed limit 296 92% 88 94% table 2. frequency of detected actions generated alerts. for manipulating the radio or the navigation system. if the driver doesn’t keep his hands on the steering wheel for a long period of time, it is considered to be a dangerous action. furthermore, if besides having no hands on the steering wheel, the driver is also looking in another direction we consider this behavior very dangerous and the system will generate an alert to notice the driver. a very important aspect taken into consideration while defining the behavioral models is the temporal evolution of an action. under normal circumstances, stretching arms (no hands on the steering wheel) or looking in another direction, these activities are performed by a driver without any serious consequences. still, in case any of these actions are performed for a longer period of time they might affect the level of concentration and of the driver’s attention. what is more, in case of an unexpected or sudden event, the driver’s speed of reaction might not be high enough, which could lead to an accident. figure 7. architecture of a digital car driver system. kinect software module was used for pose estimation. the output of the pose estimation module, consisting of the location and orientation of each body part (see figure 8) was used to assess the relation between the different body parts (e.g. arms relative to the torso, lower arms relative to upper arms). we computed the angles between the different body parts on a frame basis and also for an image sequence. the angles between head, torso and the vertical axis together with the one between head and torso provide relevant information regarding head and body position.we considered three basic orientations of the head relative to the body: straight, left, and right. the goal of our research is to test the possibility to train the basic car driver scenarios using a driving simulator and serious gaming. in table 4 we displayed figure 8. model of the upper body used for pose estimation. figure 9. hierarchy of driver activity. figure 10. incorrect detection of left arm. 89 m. toma, m. popa, l.j.m. rothkrantz acta polytechnica ctu proceedings figure 11. motion angles pattern for two actions: (a) hands on the steering wheel (b) hands raised . some of the basic scenarios. every scenario is composed of a sequence of basic actions. we selected the following basic actions: • browsing (moving position of head, looking into the mirror, inspecting the routing device or radio) • tuning (radio, mobile phone etc.) • activating (touching eyes, mouth, pulling ears etc) • selecting (putting the gear by hand in the requested position) • waving (putting left/right hand up) • driving (turning the steering wheel). we evaluated the performance of the body pose estimation module on the driving recorded data, by inspecting visually the accuracy of each detection (see figure 9 and figure 10), and we achieved 89% correct upper body detections and 58% correct upper body parts estimation. kinect proved to be very succesfull with limb detection and person tracking. another important thing about kinect’s limb tracking is that kinect is aware of occlusion and for example if one of the arms is missing, the stick configuration shows that a limb is missing. in short two main problems that are faced with kinect and the video tool are undetected limbs (see table 5) and frames with no detection at all respectively (see table 6). the detection heavily depends on the exposed action. in table 3, we show the average missing rate. the car environment can be divided into a number of regions of interest for the driver. the most important and used ones are the ’steering wheel’ (1) and the ’gears’ (2) regions, followed closely by the ’contact’ (3) and ’brake’ (4) regions. other secondary regions are the ’navigation system’ (5), the ’radio’ (6) and the ’drawer’ (7) ones. a visual representation of the defined regions can be found in figure 12. figure 12. regions of interest inside a car. we included a graphical division of the inside the car environment into regions of interest. the driver’s interaction with each object inside the car can be assessed by determining hands’ position inside a region. this information is used in combination with the motion pattern analysis in order to extract a first semantical interpretation of the possible action of the driver. movements characterized by a certain speed and amplitude, inside a specific region of interest are very likely to depict a certain type of action. for example movements having a low speed and a small amplitude associated with the ’steering wheel’, ’radio’ or ’navigation system’ regions of interest are very probable to correspond to driving, manipulating the radio or the navigation system actions. on the other hand, movements with a high speed and a large amplitude inside the ’drawer’ or the ’gears’ regions are associated with picking an item from the drawer or changing the gears actions. the rois, the transitions between the different rois, the associated type of movements, and the probable behavioural interpretation are presented in figure 13. the reasoning step was implemented using a rulebased system which received as input, features from the previously described modalities (body pose estimation, face detection, and regions of interest assessment). on a time basis the extracted features were observed and using the rules containted in the statebased model we were able to distinguish normal from potential dangerous and dangerous driving behavior. conclusion regarding the driving behavior was generated. we employed the state-based model depicted in figure 13. 90 vol. 12/2017 an intelligent co-driver surveillance system browsing tuning activating selecting waving driving browsing 23 22 55 0 0 0 tuning 0 95 5 0 0 0 activating 5 10 77 3 0 5 selecting 0 17 0 83 0 0 waving 0 0 0 14 85 0 driving 25 0 25 0 0 50 table 3. confusion matrix of selected driving actions. driver action behavior detected by kinect,video and car sensors s1 starting the car 1.check if the gear is in neutral position hand movements, movements of head (turn down) 2.start the car hand motion 3.press the throttle a little bit movements of right leg s2 driving away 1.start the car see s1 2.press the clutch to the floor movement of left leg 3.enter the gear shift in position 1 hand movement (movement of the head) 4. press the throttle pedal movement of right leg 5.release the clutch movement of left leg s3 driving away from parked position 1.start the car see s1 2.look in the inner and outer mirror if next lane is free turning the head right/left 3.switch on the left signalling light movement of hand 4.press the clutch to the floor movement of left leg 5.enter gear shift in position 1 movement of hand, movement of leg 6.press slowly the throttle pedal movement of right leg 7.turn the steering wheel to the left or right hand movement 8.turn the steering wheel to the neutral position after reaching the next lane hand movement table 4. scenarios and behavioural cues. total number of frames kinect 16% 574 video-tool 32% 574 table 5. rate of frames without detected limbs. 91 m. toma, m. popa, l.j.m. rothkrantz acta polytechnica ctu proceedings upper right arm upper left arm lower right arm lower left arm browsing 12 74 20 74 tuning 10 2 12 2 activating 2 0 2 0 selecting 12 7 12 7 waving 16 75 21 75 driving 17 20 17 20 table 6. kinect rate of frames without detected limb relative to all frames in the dataset. figure 13. state based model. we used 3 seconds as a threshold for the critical duration of a secondary action or body part orientation, such as head orientation, lower arms position, or body pose orientation. 6. conclusions in the framework of designing a digital co-driver system, we performed a study to assess the actions of a driver. based on the experiments in a car simulator it was shown that the action of the car driver can be assessed in more than 80% of the cases. but we stress the fact that the results are based on the experiments in a car simulator using students as testpersons doing their best to drive according to the rules. in real life situations the results could be quite different. we expect that in real life situations the recognition rate could be quite lower because of bad lighting conditions in the car, occlusions, position of the car driver. that is the reason that we used multiple multimodal sensors. this implies sensor fusion. to assess the driver actions we used a probabilistic rule based system. the limitation of such a system is that all possible (string of) actions have to be defined in advance. unexpected rare actions will not be detected. this is also caused by the fact that our system selects the most probable actions. in case that our system has to detect some actions with a high priority to send an alert such actions have to be labeled with a higher priority in our system. the designed system can be used as a decision support system in current cars. the car has to be equipped with some sensors and a microprocessor to process the data and generating alerts. the results can also be used in the design of self driving cars in the near future by modelling a digital car driver according to a human model. the basic positions of individual body parts, along with the motion temporal patterns and the corresponding regions of interest were fused in order to draw a conclusion regarding the driver’s possible type of activity. references [1] l. rothkrantz. surveillance angels. neural network world 24:1–25, 2014. 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[23] m. popa, l. rothkrantz. assessment of behaviour in serious games of driving simulators. international journal of intelligent games & simulation 6(2), 2011. 93 http://dx.doi.org/10.1007/978-3-319-12352-3 http://dx.doi.org/10.4236/jtts.2012.21007 http://dx.doi.org/10.1007/bf02884432 http://dx.doi.org/http://dx.doi.org/10.1016/j.ssci.2009.03.014 http://dx.doi.org/10.1109/tits.2007.895298 http://dx.doi.org/10.1109/tits.2008.2011707 http://dx.doi.org/10.1016/j.patrec.2013.01.002 acta polytechnica ctu proceedings 12:83–93, 2017 1 introduction 2 related work 3 model 4 simulation environment 5 experiments 5.1 experiment 1 detection behaviour actions 5.1.1 analysis experiment 1 5.2 experiment using kinect 6 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0089 acta polytechnica ctu proceedings 13:89–92, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app drying characteristic of rammed earth with illitic-kaolinitic clay content tereza otcovská∗, barbora mužíková, pavel padevět czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: tereza.otcovska@fsv.cvut.cz abstract. clay is a traditional construction material which has got to background with introduction of modern materials to building practice. there is no proper material available for clay constructions design due to lack of proper examination of its mechanical properties. this paper focuses on drying rate of rammed earth. drying is a primary way in which clay gets its strength. it is thus essential to know a duration needed to a construction to get dry. after pass of this time the maximum strength is attained and it is possible to load a construction. properties of unburned clay are dependent on clay mix composition. in this contribution two sets of testing bodies with different composition are presented. as a bonding agent an illitic-kaolinite clay was used. it was presupposed that amount of used clay in clay mixture has major influence on the speed of drying and a final equilibrium moisture content. this presumption was disproved. keywords: clay, rammed earth, drying rate, water coefficient, load-bearing structure. 1. introduction an unburned clay was commonly used construction material of the past. however, with the boom of new modern construction materials, unburned clay was pushed out from the business, and therefore the necessary scientific research of unburned clay mechanical properties was never performed [1–5]. these days unburned clay returns to focus again generally for three reasons. first reason is reconstruction of already built clay buildings [6–8]. second, the unburned clay constructions are compatible with the principles of sustainable development [9, 10]. third, unburned clay has a blessed impact on the internal micro-climate of the building and consequent improvement in the health of inhabitants [1, 3, 11]. one of the major methods of unburned clay manufacture is a ramming technology which this subscription concerns of. a principle of the method is ramming of earth layers into a framework. rammed earth is mostly used for load bearing constructions and for that reason is important a knowledge of the earth drying speed [1, 3]. the drying speed of unburned clay is important because drying is the only process with which unburned clay gain its strength. it is thus necessary to know drying speed to be able to estimate the time that must pass from a manufacture to full construction load [1, 3, 9, 12]. the clay mixture consists of three basic components: sand, clay and water. the sand acts as a filling agent. the clay has a function of a bonding agent similarly as a cement in concrete. and water serves for activation of bonding properties of the clay and for proper processibility of clay mixture [1, 13–15]. the clay mixtures that were experimentally examined and described in this contribution contain just these three basic components. no any additive nor additional bonding agent such as cement or lime were added to examined clay mixtures. the reason to exclude the additives was our effort to properly describe plain clay mixture. it was proved that even plain unburned clay used as a construction material features long lifespan. it is true particularly for properly designed construction details. in some cases it was discovered that by adding of another organic bonding agents the result material strength decreased. that is crucial for load-bearing construction elements. another reason why we concentrated on unburned clay without additives is a fact that addition of a cement or lime reduces positive influence of unburned clay to inner microclimate and raises environment load [3, 16, 17]. 2. test mixture and test bodies manufacture two types of clay mixtures were tested. the mixtures differed in sand to clay ratio and in amount of used water (water coefficient). the testing bodies were manufactured by ramming technology. the published results were measured minimally on six testing bodies of each mixture. the used clay mixtures are composed from a sand (see fig. 1), illitic-kaolinitic clay (see fig. 2) and water (common tap water). a sand to clay ratio and water amount were prescribed for the both clay mixtures. the behaviour of the clay mixture during drying and the speed of drying were studied in connection 89 http://dx.doi.org/10.14311/app.2017.13.0089 http://ojs.cvut.cz/ojs/index.php/app t. otcovská, b. mužíková, p. padevět acta polytechnica ctu proceedings figure 1. granularity curve. figure 2. illitic-kaolinitic clay. with clay content and water content in the mixture. the clay mixture krii contained 1/4 more clay (bonding agent) than in the mixture kriii. the mixture krii was manufactured with water coefficient 0.37. the mixture kriii was manufactured using water coefficient 0.4. the water coefficient is a ratio between weight of water and weight of clay. the actual clay mixture composition is declared in the table 1. the test bodies were rammed using electric power hammer into steel moulds of size 40×40×160 mm. only the last layer was rammed manually (see fig. 3). immediately after manufacture the testing bodies were put out of mold and weighted. after that the testing bodies were inserted into air-conditioned chamber, where constant temperature (20 °c) and relative humidity (60 %) are maintained (see fig. 4). the testing bodies were left in the air-conditioned chamber until their humidity has settled (approximately 1 month). 3. measured results the test bodies were held in the air-conditioned chamber and regularly weighted. thus the decrease of their mass was measured (see table 2, fig. 6). a frequency of weighting was the highest in the first day after manufacture (approximately once per one to three hours). during the first day after manufacture the highest drop in weight was measured (approximately 90 % from total drop of weight) (see fig. 5, table 2). figure 3. production of the test bodies. figure 4. test bodies in the air-conditioning chamber. the equation 1 was used for calculation of weight moisture content from the measured values. initial weight moisture content of the both test body sets was 9.3 %. time course of average value of weight moisture content is introduced in the table 3 and in figures 7 and 8. w = mw − md md (1) it can be seen in the figures 7 and 8 that equilibrium moisture of kriii set is 1.6 % and thus higher than moisture of krii which is 0.8 % even though that kriii set contains less clay that krii set (see table 1). this finding is surprising because it is generally believed that clay minerals bond water [1, 3, 13]. it was expected that equilibrium humidity of krii set will be higher. the reason why the equilibrium humidity is higher for clay mixture with lower clay ratio is probably that kriii set was manufactured with higher water coefficient (see fig. 9). it can be further conclude from carried on experiment that the speed of drying is independent on the amount of used clay. for the both mixtures the 90 % of moisture content (calculated from the total moisture content drop) evaporated during the first 24 hours after manufacture. equilibrium moisture content was 90 vol. 13/2017 drying characteristic of rammed earth ... set clay sand/clay ratio water/clay ratio number of test bodies krii illite-kaolinitic 75/25 0.37 6 kriii illite-kaolinitic 80/20 0.4 6 table 1. composition of the clay mixture batches and number of test bodies. figure 5. time course of the test bodies weight after the first 30 hours from manufacture. figure 6. time course of the test bodies average weight. reached after approximately 25 days. the form coefficient of test bodies defined by equation 2 was 0.8889. cs = vtb atb (2) 4. conclusion the carried out experiment showed that the amount of used water in mixture can have more substantial influence to resulting equilibrium humidity than amount of used clay. however, it is necessary to verify this finding and state the actual influence in detail. it should be acomplished by testing another sets of clay mixtures that would be manufactured in wider range of sand/clay ratio and water/clay ratio. it is also necessary to focus on microstructure of tested material to reliably describe influence of water and clay ratios on unburned clay drying courses. figure 7. time course of weight moisture content. figure 8. time course of weight moisture content – logarithmic scale of time figure 9. ratio of the clay and water in mixture and resultant equilibrium humidity. 91 t. otcovská, b. mužíková, p. padevět acta polytechnica ctu proceedings set time elapsed after manufacture [hour] 0 0.7 1.7 3 4 7 22 24 28 31 48 144 157 192 528 709 krii 592 582 564 551 550 547 547 kriii 588 587 583 575 551 550 549 548 548 547 table 2. time course of the test bodies average weight [g]. set time elapsed from test bodies manufacture [hour] 0 0.7 1.7 3 4 7 22 24 28 31 48 144 157 192 528 709 krii 9.3 7.3 4.1 1.6 1.3 0.8 0.8 kriii 9.3 9.1 8.3 6.9 2.4 2.2 1.9 1.7 1.7 1.6 table 3. time course of the test bodies average weight moisture content [%]. the obtained results also indicate that the drying speed is not dependent on the amount of the clay in a clay mixture. for test bodies with form coefficient of 0.8889 the drying was almost complete (reaching of equilibrium humidity) after the first 24 hours. list of symbols w weight moisture content [–] mm weight of damp sample [g] md weight of dry sample [g] cs form coefficient [m] vtb volume of test body [m3] atb surface of test body [m2] acknowledgements the financial support of this work by the faculty of civil engineering, czech technical university in prague (sgs project no. sgs16/201/ohk1/3t/11) is gratefully acknowledged. we would like to express our thanks to lb minerals, s.r.o. company for free supply of material necessary for 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[17] i. zabalza bribián, a. valero capilla, a. aranda usón. life cycle assessment of building materials: comparative analysis of energy and environmental impacts and evaluation of the eco-efficiency improvement potential. building and environment 46(5):1133–1140, 2011. doi:10.1016/j.buildenv.2010.12.002. 92 http://dx.doi.org/10.1016/j.conbuildmat.2013.12.067 http://dx.doi.org/10.1016/j.buildenv.2008.07.001 http://dx.doi.org/10.1016/j.buildenv.2010.12.002 acta polytechnica ctu proceedings 13:90–93, 2017 1 introduction 2 test mixture and test bodies manufacture 3 measured results 4 conclusion list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0038 acta polytechnica ctu proceedings 7:38–42, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app uhpc sandwich structures with composite coating under compressive load jan markowski∗, ludger lohaus institut für baustoffe, leibniz universität hannover, appelstraße 9a, 30167 hannover, germany ∗ corresponding author: j.markowski@baustoff.uni-hannover.de abstract. ultra-high-performance concrete (uhpc) sandwich structures with composite coating serve as multipurpose load-bearing elements. the uhpc’s extraordinary compressive strength is used in a multi-material construction element, while issues regarding the concrete’s brittle failure behaviour are properly addressed. a hollow section concrete core is covered by two steel tubes. the outer steel tube is wrapped in a composite material. by this design, uhpc is used in a materialand shape-optimised way with a low dead weight ratio[1] concerning the load-bearing capacity and stability[2]. the cross-section’s hollow shape optimises the construction’s buckling stability while saving self-weight. the composite coating on the column’s outside functions both as a layer increasing the construction’s durability and as a structural component increasing the maximum and the residual load capacity. investigations on the construction’s structural behaviour were performed. keywords: uhpc, sandwich structure, hollow tubular cross-section, composite materials, lightweight. 1. introduction the compressive strength of an ultra-highperformance concrete (uhpc) (fc) exceeds 130 n/mm2. concrete with compressive strengths of 200 n/mm2 can be unerringly produced. one of the challenges to face while working with uhpc is its brittle failure behaviour. since abrupt failure of building components is to be avoided, constructive solutions have to be found. this article addresses the design and mechanical system (see para.2.1) and the production (see para.2.3.2) of uhpc sandwich structures with composite coating, while the individual materials are described in para.2.2. the uhpc sandwich structures with composite coating exploit the concrete’s properties and provide ductile structure behaviour, as proven in experiments (see paragraph 2.3). a test set-up (see para.2.3.3) and results (see para.2.3.4) of axial compressive tests of specimens with and without composite coating are described. the results are discussed in para.2.3.4. 2. uhpc sandwich structures with composite coating as shown in figure 1, uhpc sandwich structures with composite coating consist of four layers of different materials. a hollow section uhpc core is covered by two concentrically arranged steel tubes which can be understood as lost framework and reinforcement of the uhpc core[3]. the outer steel tube is wrapped in a composite material layer. the composite layer fulfills different tasks: its fibres act as a structural component (as described in para.2.1) and its matrix protects the outer steel tube’s surface from corrosion. figure 1. uhpc sandwich structure with composite coating, cross-section. 2.1. mechanical system – conceptual approach uhpc sandwich structures with composite coating serve as multipurpose load-bearing elements, predominantly for compressive loads. axial loads are borne especially by the concrete’s cross-sectional area. the inner and outer steel tube are ideally not involved in the axial load-bearing (as far as the serviceability limit state is not exceeded), nor is the composite material layer. pure uhpc has the most favourable ratio of dead weight to compressive strength in comparison to steel. in order to increase the second moment of area and, thereby, the cross-section’s buckling resistance, a hollow cross-section design is chosen. despite the uhpc’s brittle failure behaviour, a ductile behaviour of the construction can be achieved by the interaction of the uhpc core and the construction’s other layers. after exceeding the ultimate limit 38 http://dx.doi.org/10.14311/app.2017.7.0038 http://ojs.cvut.cz/ojs/index.php/app vol. 7/2017 uhpc sandwich state, the construction is still able to bear further loads by the lateral supporting effects of the steel tubes and the composite layer. the composite material’s fibres are arranged in peripheral direction. consequently, no axial loads are borne by the composite material while buckling and detachment issues are avoided as confinement effects increase both the construction’s load-bearing capacities and its ductility. longitudinal strains of the columns under axial compression result in transverse strains. while the steel’s poisson’s ratio is larger than the concrete’s poisson’s ratio, peripheral tensile stresses occur in the concrete core. radial expansions can be converted into peripheral stresses by the composite material’s fibres, reducing the harmful tensile stresses inside the concrete core and, therefore, multiaxial compressive states of stress can be activated. due to these unidirectional properties of the composite material wrapping, the outer steel tube can be chosen with a relatively low thickness, resulting in further weight savings compared to conservative constructions. 2.2. materials 2.2.1. steel non-alloy quality steel, grade dc01, is used for the outer steel tube and conventional structural steel s235 is used for the inner steel tube. both materials have a characteristic yield stress (fy) of 235 n/mm2. 2.2.2. ultra-high performance concrete the concrete’s mix design and its basic material parameters are shown in table 1 and 2. component type kg/m3 cement cem i 52.5r hs/na 175 microsilica uncompacted 825 fine quartz powder 50/16 µm 200 sand 0.125/0.5 mm 975 water 185 superplasticizer polycarboxylate 28 w/c = 0.24 water to cement ratio w/b = 0.20 water to binder ratio table 1. the design of the concrete mix used. 2.2.3. composite material the carbon fibre reinforced plastic composite consists of • a unidirectional carbon band, which is aligned in the peripheral direction of the specimen (see fig. 2) and • a matrix of epoxy resins, which is predominantly used for rotor blades for wind turbines and shipbuilding. average density 2.323 kg/dm3 slumpflow (sf) (1) 900 mm compressive strength (2,3) 220.0 n/mm2 flexural strength(2,4) 13.3 n/mm2 (1) according to din en 12350-8[4] (2) after thermal treatment at 90 ◦c for 48 h, started 24 h after casting, cube 100 × 100 mm (3) according to din en 12390-3[5] (4) according to din en 196-1[6] table 2. the basic material parameters of the concrete used. carbon band material carbon fibre (ht) width 25.0 mm thickness 0.3 mm weight 350.0 g/m2 yarn count 400.0 tex matrix epoxy resin hexion epikote epoxy hardener hexion epikure heat-curing 12 h, 70 ◦c table 3. the composite material used. figure 2. unidirectional carbon fibre band 2.3. experimental research in order to demonstrate the composite material layer’s impact on the mechanical system, the maximum loadbearing capacities of specimens with fibre reinforced plastic composite coating (series b, two specimens) are compared to those of specimens without composite coating (series a, three specimens). 2.3.1. specimen geometry the specimen’s properties are shown in table 4. 2.3.2. production of the specimens in a first step, the steel tubes are aligned perfectly concentrically. afterwards, the concrete is placed by means of a tremie pipe (contractor method). after heat curing the concrete (see para.2.2.2), the composite material is applied onto specimen b in a hand lay-up process by means of a mounting fixture that allows rotation around the longitudinal axis. in a last step, the composite material is heat cured. a specimen of series b is shown in figure 3. 39 jan markowski, ludger lohaus acta polytechnica ctu proceedings series a: without composite layer inner steel tube: ∅outside = 133.0 mm ts, inner = 2.9 mm outer steel tube: ∅outside = 169.0 mm ts,outer = 1.0 mm concrete core: ∅mid−line = 155.0 mm tc = 19.2 mm length: l = 500 mm weight: m = 15.39 kg series b: with composite layer composite material: 8 layers of carbon band ≈ 3.0 mm (inc. matrix) weight: m = 16.64 kg in all dimensions as series a table 4. properties of the two series figure 3. specimen of series b 2.3.3. test set-up the experimental work focuses on the maximum loadbearing capacity and the ductility. the specimens are loaded in a displacement-controlled (load speed 0.25 mm/min) test by a hydraulic cylinder to failure and beyond. an integrated calotte levels possible tilting. normal force and displacement are measured. in order to introduce the load predominantly into the concrete core and to avoid cracks in the load introduction area, the specimens are grouted in a load introduction panel. the same concrete mix design as used for the specimen is used as grout material. figure 4 shows the test set-up containing the specimen, the load introduction panels, three inductive displacement transducers and protective equipment. 2.3.4. results table 5 shows the results for all specimens of series a and b. force–distance diagrams for two selected exfigure 4. experimental set-up figure 5. force–distance diagrams, two selected examples fn,max fn,residual series a, specimen 1 1, 395 kn 554 kn series a, specimen 2 1, 389 kn 675 kn series a, specimen 3 1, 260 kn 684 kn series a average 1, 348 kn 638 kn series b, specimen 2 1, 774 kn 1, 610 kn series b, specimen 2 2, 073 kn 950 kn series b average 1, 924 kn 1, 280 kn table 5. results of the experimental research 40 vol. 7/2017 uhpc sandwich amples are presented in figure 5. positive fn -values represent compressive forces for means of presentation. figure 6 shows the two specimen after testing. the broken concrete core of the specimen shown causes a buckling formation on the outer steel tube. the inner steel tube (not shown in fig. 6) is also affected. the buckling appeared when the maximum load capacity of 1, 260 kn was reached. the specimen of series b shown appeared to be undamaged from the outside until a displacement of 10.0 mm is reached. at this point, only single sectors of the composite material failed to a limited extent. when the displacement reached 10.0 mm, a larger amount of fibres failed, causing an extensive detachment in the specimen’s middle section. under the composite layer, the outer steel tubes showed buckling formations that are comparable to those of specimen a. figure 6. series a specimen 3 (left) and series b specimen 1 (right) after testing 3. discussion as the results show, the maximum load-bearing capacity and the residual load-bearing capacity of uhpc sandwich structures can be increased by the addition of a composite layer. the ductility (in this context, the ratio of the residual force to the maximum force) is affected as well, but the results differ. while the ductility of specimen 1 of series b can be increased significantly (the residual force is at 90 % of the ultimate load), the residual force of specimen 2 is at 59 % of its ultimate load. it is still higher than the average ductility of the specimens of series a (47 %), but, in view of the large variation, the effect is negligible. it is assumed that different modes of failure are the reason for this variation. while the concrete core of specimen 1 fails first and is still supported laterally by the composite material, it is suggested that in the case of specimen 2, the fibres fail first and lead to a sudden drop in the load-bearing capacity (from 2, 073 to 950 kn) bearing capacity. table 6 shows a simple approach to estimate the construction’s load-bearing capacity. the composite layer is not taken into account because of its peripheral fibre orientation. cross-sectional area (specimen a and b) steel tubes: as = 1, 713 mm2 concrete core: ac = 7, 959 mm2 load-bearing shares (simple assumption): fn,s = as ∗ fy = 402.4 kn = 16.7 % ∗ fn,total fn,c = ac ∗ fc = 1, 750.9 kn = 81.3 % ∗ fn,total fn,total = fn,s + fn,c = 2, 153.3 kn load-bearing capacity (experimental results): specimen a: fn,average = 1, 348 kn specimen b: fn,average = 1, 924 kn table 6. maximum load-bearing capacity estimation, simple approach the construction’s maximum load-bearing capacity is overestimated by the simple approach. the reason for this could be • the load introduction by load introduction panels that exclude loads partially from the steel tubes; • effects caused by stability; or/and • imperfections (material and geometry). lindschulte proposes a more sophisticated estimation model (eq. 7-1[2]) for sandwich structures without composite coating: fn = ac ∗ 0.8 ∗fc +αs−1 ∗as,−1 ∗fy +αs+1 ∗as,+1 ∗fy (1) where is as,−1 = area of the inner steel tubes cross-section as,+1 = area of the outer steel tubes cross-section αs−1 = stability reduction factor concerning to χ(λ) according to eurocode 3, inner steel tube αs+1 = stability reduction factor concerning to χ(λ) according to eurocode 3, outer steel tube applying lindschulte’s model to the series’ geometry shown leads to fn = 1, 400 kn assuming a non-load-bearing role of the steel tubes. this estimation fits the experimental results of series a (fn,average = 1, 348 kn) well. the increase 41 jan markowski, ludger lohaus acta polytechnica ctu proceedings of the maximum load-bearing capacity of series b cannot be explained with this model. confinement effects as a result of the composite layer have to be considered as a reason for this. the composite material’s fibres bear tensile stresses in a peripheral direction, that are caused by transversal strains in a radial direction under axial compression. without these fibres, these radial strains lead to tensile stresses in the concrete-core. due to the confinement effect, all compressive stress states in the concrete core are reached, leading to a higher maximum loadbearing capacity. both specimen of series b also have a higher residual load-bearing capacity. this can be explained by the lateral supporting effect of the composite material. after the concrete is broken, it is kept in place by the inner steel tube and the outer steel tube, which is additionally supported by the composite material. 4. outlook uhpc sandwich structures with composite coating could be beneficially used, for example, in large frameworks of offshore megastructures[7]. in order to prove the concept, further investigations have to be performed. the following subjects require special attention: • the identification of different failure modes. • numerical simulations to prove the mechanical behaviour assumed. • further experimental research with different geometries in order to derive a model for the prediction of the maximal load and the residual load. • the component’s behaviour under pure bending and bending with normal force. • the component’s behaviour under cyclic load. • junctions and load introduction. list of symbols ε strain [–] ac concrete area [mm2] as steel area [mm2] fn maximum normal force [kn] fn,c normal force, concrete’s share [kn] fn,s normal force, steel tubes’ share [kn] fc concrete’s compressive strength [n mm−2] fy steel’s yield stress [n mm−2] u displacement [mm] w/c water to concrete ratio [–] w/b water to binder ratio [–] acknowledgements the research project is supported by the german research foundation (dfg) within the scope of priority programme 1542 "lightweight building with concrete". the authors would like to express their gratitude for the financial support. references [1] n. scholle, l. lohaus, n. lindschulte. weight-saving potential of hybrid tube structures. proceedings of hipermat 2016, 4th international symposium on ultra-high performance concrete and high performance construction materials 2016. [2] n. lindschulte. druckverhalten von rohren aus ultrahochfestem beton mit stahlblechummantelung. institut für baustoffe, leibniz universität hannover, appelstr. 9a, 30167 hannover, de, 1st edn., 2013. [3] n. scholle, l. lohaus. hybrid tubes: material-adapted construction with ultra-high performance concrete. iass2015 annual international symposium on future visions 2015. [4] din en 12350-8: testing fresh concrete part 8: self-compacting concrete slump-flow test; german version en 12350-8:2010. [5] din en 12390-3: testing hardened concrete part 3: compressive strength of test specimens; german version en 12390-3:2009. [6] din en 196-1: methods of testing cement part 1: determination of strength; german version en 196-1:2005. [7] n. scholle, l. lohaus. uhpc-steel hybrid tube components for application in offshore support structures. the annual international offshore and polar engineering conference (isope) 2016. 42 acta polytechnica ctu proceedings 7:38–42, 2017 1 introduction 2 uhpc sandwich structures with composite coating 2.1 mechanical system – conceptual approach 2.2 materials 2.2.1 steel 2.2.2 ultra-high performance concrete 2.2.3 composite material 2.3 experimental research 2.3.1 specimen geometry 2.3.2 production of the specimens 2.3.3 test set-up 2.3.4 results 3 discussion 4 outlook list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0121 acta polytechnica ctu proceedings 13:121–124, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app coupled simulation for fire-exposed structures using cfd and thermo-mechanical models stanislav šulca, ∗, vít šmilauera, františek waldb a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b czech technical university in prague, faculty of civil engineering, department of steel and timber structures, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: stanislav.sulc@fsv.cvut.cz abstract. fire resistance of buildings is based on fire tests in furnaces with gas burners. however, the tests are very expensive and time consuming. this article presents a coupled simulation of an element loaded by a force and a fire loading. the simulation solves a weakly-coupled problem, consisting of fluid dynamics, heat transfer and mechanical model. the temperature field from the computational fluid dynamics simulation (cfd) creates cauchy and radiative boundary conditions for the thermal model. then, the temperature field from element is passed to the mechanical model, which induces thermal strain and modifies material parameters. the fluid dynamics is computed with fire dynamics simulator and the thermo-mechanical task is solved in oofem. both softwares are interconnected with mupif python library, which allows smooth data transfer across the different meshes, orchestrating simulations in particular codes, exporting results to the vtk formats and distributed computing. keywords: thermo-mechanical model, virtual furnace, computational fluid dynamics, multiphysics, fire resistance. 1. introduction several single physical models exist for description of particular physical phenomena. multi-physical coupled problems extend applications to more complicated tasks. in order to alleviate coupling among models, we created a python library which integrates solutions for computing simulations of structures exposed to fire. this paper presents approach how to interconnect single physical models. 1.1. fire dynamics analysis the issue of predicting fire process in a defined space domain is solved with the fire dynamics simulator software (fds) [1], developed at nist. we modified the structure of the code to be able to call each computational time-step individually, we added several interface functions to get the dimensions of meshes, the temperatures in nodes, etc. 1.2. thermo-mechanical analysis the thermo-mechanical analysis of the element takes place in oofem [2, 3] software. the thermal analysis uses a temperature field from fds, which defines cauchy and radiative boundary conditions on the element. then, the solution of thermal analysis passes temperature to the mechanical analysis, which induces the thermal strain and modifies the material parameters as yield strength and elastic modulus. temperature will also be used for timber structures where it controls carbonization process and changes the thermal conductivity. 1.3. adiabatic surface temperature the heat transfer to the construction is realized by convection and radiation. fds can compute a quantity called "adiabatic surface temperature" (ast), which includes both convection and radiation influence. we compute the total heat flux qtot into the element with a method by wickstrom [4] from the following equation: qtot = εσ ( t 4ast − t 4) + h (tast − t) (1) where ε is the emissivity, σ is the stefan-boltzmann constant and h is the heat transfer coefficient. it allows us to realize the complicated task of radiation transfer very easily. 2. interfacing fds and oofem using mupif interfacing fds and oofem using mupif [5, 6] library, which provides several useful tools such as exporting data to vtu format, parallel computations and advanced handling with data fields and meshes. both the fds fortran and oofem c++ codes provide application interface for communication and they are compiled as shared libraries. top-level steering script controls execution of both codes, data synchronization etc. the flowchart of the codes is given in fig. 1. the fds task consists of one or several uniform hexahedral meshes. in the beginning of the computation the geometries of all meshes are mapped from 121 http://dx.doi.org/10.14311/app.2017.13.0121 http://ojs.cvut.cz/ojs/index.php/app s. šulc, v. šmilauer, f. wald acta polytechnica ctu proceedings figure 1. flowchart of the codes. figure 2. data workflow. fds to mupif. the meshes contain discretized temperature field. we also load the coordinates of the points in which ast will be computed and create a mesh analogically. the only difference here is that the mesh consists only of points and has no elements. in this case, oofem uses no interpolation and uses a value of the closest point. we note that the ast points should not be exactly on the element surface, they must be in the air close to the surface. here we describe the loop over one oofem timestep, see the data workflow in fig. 2 and the flowchart in fig. 1. first, fds is called to process several computational steps until we reach the length of our defined time-step of the thermo-mechanical task. the fds application determines its time-step itself and it is usually much shorter than the oofem time-step. then, we load the temperature fields from fds and copy them into the mupif fields. the field of ast values is copied into the oofem boundary-condition field. then, oofem is asked to solve computational step of the thermal task and passes the computed temperatures of the construction to the mechanical task. the material parameters are updated due to the change of the temperature and the temperature strain is added. the mechanical response of the construction is obtained, which is the end of the computational loop. as the fds analysis is very time-consuming process and the computations can last couple of days, we use another mupif functionality to store the meshes with all the temperatures in all defined time-steps in files. in this regard, it is possible to precompute fds simulation and compute multiple thermo-mechanical analyses separately and much faster. coarse fds mesh can be refined in areas where structure elements occur. this improves precision in temperature predictions and gives more realistic results of the thermo-mechanical task. the mupif library has a tool to export the loaded fds meshes to vtu, so that the fds temperature fields can be exported into this useful format. oofem exports its results to vtu format as well. 3. example we demonstrate functionality of coupled cfd-thermomechanical model on a prototypical task, consisting of a steel beam placed into a furnace, see the geometry in fig. 3. the furnace consists of two meshes. the bigger one covers the whole furnace and has 1000 cells, while the second one around the beam has 4800 cells. the âůźmesh structure is displayed in the results in fig. 10. the inner space of the furnace is a cube with edge length 0.8 m. the furnace is heated by 1 natural gas burners, placed in the middle of its floor. the output of the burner is 400 kw during the whole simulation. in the middle of the ceiling is a chimney. both the holes are square-shaped with edge 122 vol. 13/2017 figure 3. furnace geometry. figure 4. cross-section of the beam. size 0.2 m. the simply-supported beam is made from steel i profile, see the geometry of its cross-section in fig. 4. it is placed 0.52 m from the floor, horizontally in the centre. it is loaded only by its weight and the temperature. this example aims to the thermal task, because the used mechanical material model has constant parameters, non-dependent on temperature. the material parameters are e = 200 gpa, ν = 0.3, ρ = 7850 kg/m3 and α = 12 × 10−6. the length of the beam is 0.5 m. steel conductivity is displayed in fig. 6 and capacity is displayed in in fig. 7. emissivity of the steel surface is 0.8 and the heat-transfer coefficient is 20 w.m−2.k−1. the oofem model of the beam consists of 540 cells, which is enough for our purpose, see fig 5. the results were computed end exported to vtu files. fig. 10 shows a cross-section through the middle of the furnace. we can see the two fds meshes, one coarse on the whole furnace and one finer around the beam. detailed temperatures of the beam are displayed in fig. 11, the stress magnitude σmax is in fig. 13. figure 5. oofem mesh of both thermal and mechanical task. figure 6. steel conductivity. figure 7. steel capacity. figure 8. thermal conductivity at t = 480 s. figure 9. thermal capacity at t = 480 s. 123 s. šulc, v. šmilauer, f. wald acta polytechnica ctu proceedings figure 10. temperatures at t = 480 s, section in the middle of the furnace. figure 11. temperature at t = 480 s. figure 12. stress at t = 480 s. figure 13. displacement z at t = 480 s. 4. conclusions this paper presents an approach for computing weakly-coupled problems, focusing on computational fluid dynamics coupled with thermo-mechanical analysis. temperature field from fds was validated previously [7], performance of structure elements presents a consecutive step. validation of this multi-physical approach happens during 2017 on steel and timber and concrete structural elements. acknowledgements we gratefully acknowledge the financial support from the czech technical university in prague, the grant sgs17/042/ohk1/1t/11. we also gratefully acknowledge the financial support from the czech science foundation, the grant gacr 16-18448s. references [1] nist. fds, fire dynamics simulator, project home page, 2017. http://www.pages.nist.gov/fds-smv/. [2] b. patzák. oofem project home page, 2000. http://www.oofem.org/. [3] b. patzák. oofem – an object-oriented simulation tool for advanced modeling of materials and structures. acta polytechnica 52(6):59–66, 2012. [4] u. wickstrom, d. duthinh, k. mcgrattan. adiabatic surface temperature for calculating heat transfer to fire exposed structures. in proceedings of the eleventh international interflam conferrence, interscience communications. london, 2007. [5] b. patzák. mupif project home page, 2010. http://www.mech.fsv.cvut.cz/mupif/. [6] b. patzák, v. smilauer, g. pacquaunt. design of a multiscale modelling platform. in 15th international conference on civil, structural, and environmental engineering computing. prague, czech republic, 2015. [7] k. cábová, v. asgf, g. sdfggf. modelling of standard fire test. in engineering mechanics. svratka, czech republic, 2017. 124 http://www.pages.nist.gov/fds-smv/ http://www.oofem.org/ http://www.mech.fsv.cvut.cz/mupif/ acta polytechnica ctu proceedings 13:122–125, 2017 1 introduction 1.1 fire dynamics analysis 1.2 thermo-mechanical analysis 1.3 adiabatic surface temperature 2 interfacing fds and oofem using mupif 3 example 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0020 acta polytechnica ctu proceedings 15:20–24, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app cement matrix containing micronized recycled concrete and the influence of additives jakub ďurejea, ∗, zdeněk prošeka, b, jaromír hrůzaa, pavel tesáreka a ctu in prague, faculty of civil engineering, thakurova 7, 166 29 prague 6, czech republic b ctu in prague, uceeb, trinecka 1024, 273 43 bustehrad, czech republic ∗ corresponding author: jakub.dureje@fsv.cvut.cz abstract. the article deals with the selection of a suitable additives in cement composite material for the production of a masonry block. the blocks contains recycled concrete which replacing 40% cement. the recyclate comes from concrete drainage gutters. based on previous measurements, two hardening accelerators and one superplasticizer were selected. to determine the effect of these additives on the recycled cement composite, pairs of samples sets were produced, of which one set contains concrete recyclate and the other one does not contain it. the dynamic modulus of elasticity and the dynamic shear modulus were measured by the resonance method at the age of samples 1, 7, 14, 21 and 28 days, and by an ultrasound method at 28 days of age. then was measured compressive strength. all values were compared with the values measured on the reference samples. keywords: recycled concrete, micro-milled concrete, influence of additives, resonance method, dynamic modulus of elasticity, dynamic shear modulus. 1. introduction concrete waste ,which comes from demolition of buildings, can be recycled but there can be problem with different properties and inhomogeneity of this waste due to the presence of other materials (bricks, wood, glass, reinforcement, etc.). the quality of the recycled concrete depends on the care of sorting this old waste. unfortunately, waste sorting is negatively reflected in the price of this raw material. recycled concrete is most commonly used as the underlayer under the road. usually is used fraction of recyclate greater than 1 mm, the use of finer fractions brings many complications and additional financial costs [1, 2]. the article deals with the measurement of the dynamic modulus of elasticity and compressive strength of samples containing cement, various additives and concrete recyclate. recycled concrete was used, including the finest 0-1mm fractions, which were then finely grinded using a highspeed milling process. the use of this fine fraction, which has almost no use today, is an advantage of the material. these fine fractions of concrete recyclates are often stored in landfills, and their use can reduce the volume of stored waste materials which is also environmental friendly. finely ground recycled concrete was used in cement composites as a filler with certain bonding properties. from previous experiments, it has been shown that with the increasing percentage of recyclate in the cementitious composite, compressive strength decreases [3, 4]. in order to reach a certain compressive strength, various additives to concrete has been used which cause an increase the initial and final compressive strengths. the examined material is planned to be used for the production of concrete blocks for the construction of single-storey buildings in the middle east. concrete blocks could then be lightened to achieve thermal insulation properties and then used as a perimeter masonry. 2. materials and specimens a total of 8 sets of samples were produced, each set containing 6 specimens. the first half of the samples were produced with no recycled concrete in the second half was replaced 40 % of cement by recycled concrete. the two sets were reference (with and without recyclate), the other sets contained additives. for the production of samples was used cem i 42.5 r portland cement radotin from českomoravsk y cement. recycled concrete comes from drainage gutter. a fraction 0/16 mm was used for high-speed milling process. the resulting grain size was <0.25 mm. the fineness of the ready-to-use recycled concrete approached the fineness of the cement with a surface area 361 m2. based on previous measurements, suitable additives were selected. superplasticizing additive (p3) based on modified polycarboxylates which is especially designed for concrete with requirements for high short and final compressive strengths. the accelerator (u1) which does not impair the workability of the fresh mixture, slightly reduces the amount of mixing water. it increases the initial and final strength of the concrete. the (u2) accelerator is composed of soluble mineral salts and organic compounds. combination of this substances has effect which accelerating hardening of concrete (table 1). the water ratio (weight of water divided by mass of cement and recyclate) w / (c + r) was determined on the basis of the flow expansion test after 10 and 20 impulses to achieve approximately equal expansion of the fresh concrete 20 http://dx.doi.org/10.14311/app.2018.15.0020 http://ojs.cvut.cz/ojs/index.php/app vol. 15/2018 cement matrix containing micronized recycled concrete... mixture for all mixtures. the water ratio ranged from 0.21 to 0.31, more mixed water was used for samples containing recycled concrete to achieve approximately the same workability (table 2). laboratory molds measuring approximately 40 × 40 × 160 mm were used to make samples. test specimens were removed from the molds after 24 hours and stored under water in a laboratory environment at a temperature of 21 ± 2 ◦c and a relative humidity of 50 ± 5% for 28 days. 3. experimental methods for all samples, the dynamic modulus of elasticity at the age of 1, 7, 14, 21 and 28 days was continuously measured. a resonance method was used to measure the dynamic modulus of elasticity of samples. advantage of this method is a speed of measurement and repeatability because it is non-destructive method. the method is based on the calculation of dynamic modulus of elasticity and dynamic shear modulus based on measured fundamental resonance frequencies of longitudinal, transverse (bending) and torsional oscillation. the resonance frequencies were measured in the laboratory using the brüel & kjær measuring system, consisting of the brüel & kjær type 4519-003 response sensor, the brüel & kjær type 8206 impact hammer, the brüel & kjær fron-end 3560b-120 measuring center and the control notebook. the sample was placed on supports of sufficiently low stiffness so that the supports were placed at the nodal points of the first waveform corresponding to the given oscillation longitudinal, transverse or rotational (fig. 1). the response sensor (s) was attached to the sample to the appropriate position and the oscillation was excited by the impact hammer (b). both signals, response and excitation were recorded and converted to a frequency domain using a fast fourier transform. subsequently, the pulse labshop software version 14.0.1 has calculated the transmission function. based on the resonance peaks in the transmission function, the fundamental resonance frequency of the given oscillation was determined. from the measured resonance frequency, weight, and sample dimensions, the dynamic modulus of elasticity and shear modulus were determined [5]. the dynamic modulus of elasticity was controlly measured by the ultrasonic pulse method at the age of 28 days. the method is based on the principle of measuring the velocity of ultrasound pulse in the material under investigation. ultrasound pulse spreads through each material at a different speed, it depends on its composition and its properties. implus is induced by an electroacoustic exciter attached on the surface of the sample to be tested. when the impulse passes through a path of known length (sample length), the vibration pulse is converted to an electrical signal by the sensor, and the electronic time circuit allows to measure the pulse passage time [6]. measurement was carried out by attaching sensors to the opposite sides of the sample, and then measuring the time of ultrasound wave spreading through the length of the sample. the sonic vita gel was used to achieve perfect contact between the specimen and the ultrasonic sensor. to calculate the dynamic modulus of elasticity by the ultrasonic method, the following equation was used: eult = ρ. l2 t2 . 1 k2 .10−9 (1) eult the modulus of elasticity measured by the ultrasonic method [gpa], ρ is the bulk density of the measured sample [kg·m−3], l is the length of the measured sample [m], t is the time of ultrasound wave spreading through length of the sample [s], k is the coefficient of the dimensionality [-]. subsequent destructive tests were performed on the samples. compressive strength was tested in an ev 40 hydraulic press with a maximum force of 400 kn according to en 12390-3 testing of hardened concrete part 3: compressive strength of specimens [7]. the hydraulic press spread pressure over the area of the 40 × 40 mm steel plates to the test sample until it was crushed. load speed was 0.3 mm/s. 4. results and discussions dynamic modulus of elasticity of each set are mean from all of 6 samples in the set. control measurements and weighing have been performed on samples before every measuring by resonance method. the evaluation was performed on the basis of measured frequencies from longitudinal oscillation (figure 2). to verify the accuracy of the results, the values of the dynamic modulus of elasticity from the transverse oscillations, which are almost identical to the values of the longitudinal oscillation modules, are also evaluated, which confirms the accuracy of the measurement. from the results of the dynamic modulus edyn, it is clear that the maximum increase in edyn is during the first 7 days. between 7 and 28 days there are no significant changes to edyn. the highest values of edyn reach samples with superplasticizer p3, in comparison with samples without recyclate (cp3 23.0 gpa) and in comparison with samples with recyclate (crp3 19.0 gpa). the difference between the cp3 and crp3 sets is approximately 4.0 gpa. the drop in values of about 4 gpa is noticeable for all pairs of ingredients with a given additive and without or with recycled concrete. we can declare that 40 wt. % of recycled material causes the dynamic elastic modulus to drop by about 4 gpa in 28 days due to less matrix coherence. unsatisfactory values reached samples with hardening accelerators (cu1, cu2, cru1, cru2) whose edyn values are only slightly higher or even lower than the edyn reference sets (c, cr). this additives cause initial increases in strength, therefore at the age of 7 days, the samples with recycled material reached higher values than the reference mixture. however, a subsequent slower increase in the modulus 21 j. ďureje, z. prošek, j. hrůza, p. tesárek acta polytechnica ctu proceedings designation product based on normal dosing density ph [wt. % of cement] [kg/m3] [-] p3 modified polycarboxylates 1.0 1060 ± 10 6.5 ± 2.0 u1 unchlorinated liquid 1.0 − 3.0 1450 ± 30 7.5 ± 2.0 u2 soluble mineral salts and 1.5 1300 ± 10 9.5 ± 1.0 organic substances table 1. selected additives and its properties. micronized set/ material cement concrete recyclate water ratio additive [g] [g] [-] [wt. % of cem. + recycl.] c 1000 0 0.29 0.0 cp3 1000 0 0.21 1.0 cu1 1000 0 0.29 1.5 cu2 1000 0 0.29 1.5 cr 600 400 0.31 0.0 crp3 600 400 0.25 1.0 cru1 600 400 0.30 1.5 cru2 600 400 0.30 1.5 table 2. composition of the samples. of elasticity of mixtures with hardening accelerators resulted in some cases (cu1, cru1) that modulus not reaching the value of reference (c, cr). the dynamic shear modulus is also affected by replacing cement with recycled concrete. in each of the two groups (with and without recyclate), samples with the superplasticizer p3 reached highest values of the gdyn (cp3 9.3 gpa, crp3 7.7 gpa). provable jump for all mixtures (figure 3) can be seen between the first day and 7 days of age when strength raises fastest in the all samples. with increasing time from 7 days onwards, the shear modulus increase is only gradual, practically negligible. negative effects on the dynamic shear module are mainly u1 and u2 solidification accelerators, especially for fine-grained recycled concrete mixtures (cru1, cru2), which achieve lower elastic moduli than reference samples. the compressive strength test was also carried out in the age of 28 days. for this test were used halfs of the beams from previous bending tests. total of 11 samples (half beams) were tested from each set. the resulting values are therefore an arithmetic mean of 11 values. the results of the test show higher values of compressive strength for mixtures without recycled concrete versus mixtures with recyclate. the greatest difference in strength between samples with and without recycled concrete is in the pair of mixtures with accelerators u1 and u2. significant reductions in the strength of cru1 and cru2 samples may be due to the adverse effects of accelerators on recycled concrete (figure 4). 5. conclusion the article deals with the selection of the most suitable additive for the production of concrete blocks containing micronized concrete recyclate. based on previous measurements, two solidification accelerators (u1 and u2) and one superplasticizer (p3) were selected [8, 9]. for comparison and identification of the influence of recycled concrete, pairs of sample sets were produced, which differed only by replacing part of the cement with micronized recycled concrete. based on results, it can be concluded that: • the highest increase in the dynamic modulus of elasticity of mixtures occurs in the first 7 days of hardening. most then within 24 hours. • 40 wt. % of recycled concrete will cause a dynamic elastic modulus drop in 28 days for all 24 samples with recycled concrete of about 4 gpa. • adding accelerators to the mixture cause slight increase of water ratio. • crp3 has proven to be the best in compression strength tests. it is the strongest tested composite with recyclate. • the u1 accelerator proved to be totally inappropriate, the value of the dynamic modulus edyn is even lower than reference. • the proposed modified crp3 mixture can be used to make lightweight concrete blocks, which can be used to masonry the perimeter walls of the singlestorey building. • the proposed modified resultant matrix (crp3 22 vol. 15/2018 cement matrix containing micronized recycled concrete... figure 1. measurement basic natural frequencies from longitudinal (left), transverse (center), and torsional (right) oscillations. figure 2. increase the dynamic modulus of elasticity over time. figure 3. increase the dynamic shear modulus over time. mixture) can be lightened because of sufficient compressive strength, eg. by foaming. list of symbols eult modulus of elasticity measured by the ultrasonic method [gpa] ρ bulk density of the measured sample [kg m−3] l length of the measured sample [m] t time of ultrasound wave spreading through length of the sample [s] re coefficient of the dimensionality 23 j. ďureje, z. prošek, j. hrůza, p. tesárek acta polytechnica ctu proceedings figure 4. compressive strength. acknowledgements this outcome was supported by the czech technical university in prague under no. sgs16/201/ohk1/3t/11, by the czech science foundation research projects gacr 17-06771s and by the ministry of education, youth and sports within national sustainability programme i, project no. lo1605. the authors also thank the center for nanotechnology in civil engineering at the faculty of civil engineering, the joint laboratory of polymer nanofiber technologies of the institute of physics, academy of science of czech republic, and the faculty of civil engineering, czech technical university in prague. references [1] a. barbudo, f. agrela, j. ayuso, et al. statistical analysis of recycled aggregates derived from different sources for sub-base applications. construction and building materials 28(1):129–138, 2012. [2] i. vegas, j. a. ibáñez, a. lisbona, et al. pre-normative research on the use of mixed recycled aggregates in unbound road sections. construction and building materials 25(5):2674–2682, 2011. [3] j. topič, z. prošek, t. plachý. influence of increasing amount of recycled concrete powder on mechanical properties of cement paste. in iop conference series: materials science and engineering, vol. 236, p. 012094. iop publishing, 2017. [4] j. hrůza. design of masonry elements with use of recycled concrete, 2016. [5] čsn 73 1372 nedestruktivní zkoušení betonu–rezonačnní metoda zkoušení betonu, 2012. [6] p. dvorský. diplomová práce využití ultrazvukové impulsové metody nástroj pro stanovení pevnosti cementů 2012. [7] čsn, en 12390-3–zkoušení ztvrdlého betonu–část 3: pevnost v tlaku zkušebních těles, 2009. [8] e. anastasiou, k. g. filikas, m. stefanidou. utilization of fine recycled aggregates in concrete with fly ash and steel slag. construction and building materials 50:154–161, 2014. [9] f. cartuxo, j. de brito, l. evangelista, et al. rheological behaviour of concrete made with fine recycled concrete aggregates–influence of the superplasticizer. construction and building materials 89:36–47, 2015. 24 acta polytechnica ctu proceedings 15:20–24, 2018 1 introduction 2 materials and specimens 3 experimental methods 4 results and discussions 5 conclusion list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0142 acta polytechnica ctu proceedings 13:142–147, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app evaluation of effective properties of basalt textile reinforced ceramic matrix composites soňa valentováa, ∗, vladimír hrbeka, b, michal šejnohaa a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b institute of theoretical and applied mechanics as cr, v.v.i., prosecká 76, 190 00 prague 9, czech republic ∗ corresponding author: sona.valentova@fsv.cvut.cz abstract. the present paper is concerned with the analysis of a ceramic matrix composite, more specifically the plain weave textile fabric composite made of basalt fibers embedded into the pyrolyzed polysiloxane matrix. attention is paid to the determination of effective elastic properties of the yarn via homogenization based on the mori-tanaka averaging scheme and the 1st order numerical homogenization method adopting a suitable representative computational model. the latter approach is then employed to simulate the response of the yarn when loaded beyond the elastic limits. the required mechanical properties of individual material phases are directly measured using nanoindentation with in-build scanning probe microscopy. applicability of the proposed computational methodology is supported by the analysis of a unidirectional fibrous composite, representing the yarn, subjected to a macroscopically uniform strain. keywords: basalt fibers, textile composites, homogenization, periodic unit cell, damage. 1. introduction the textile reinforced composites have been gaining considerable attention in many different branches of engineering. structural applications of these composites are now rather common owing to their specific strength, high stiffness, high temperature resistance, and relatively low cost in comparison to traditional materials. in addition, the possibility of influencing the resulting macroscopic properties during fabrication makes them very interesting. on the other hand, predicting the effective elastic and fracture properties and thus estimating the real behavior under variety of loads calls for a quite complicated computational analysis. such an analysis grounds on the knowledge of properties of individual components, complex inner geometry depending on a technology of composite fabrication, and on a type of textile reinforcement. mostly the carbon and glass fibers as a reinforcement are used, but attention is also accorded to the basalt fibers especially due to their good thermal stability, important in the manufactured process using high temperatures when ceramic matrices are needed in the final product. they can also substitute the carbon fibers in most applications at a fraction of price. this paper deals with the evaluation of effective elastic properties of the basalt textile reinforced ceramic matrix composite at the level of yarns (microscale) using the mori-tanaka scheme and the finite element method, predicting its strength and characterizing the failure behavior. herein, a two-dimensional plain weave textile is examined. first, experimental measurements are carried out to determine properties of individual components. next, a suitable geometrical model based on the statistically equivalent periodic unit cell is designed and finally the nonlinear computational analysis is realized. 2. experimental measurements of the material parameters a specimen of basalt fiber ceramic composite was embedded into an epoxy resin and left for 24 hours to harden in a cylindrical mold (25 mm in diameter). approximately a 15 mm thick part of the sample was separated using a diamond cut-off wheel. to ensure an adequate surface roughness, the investigated sample was ground on an md-piano plate using 1200, 2000 and 4000 grain cm−2 grid and water as a lubricant under 5 n force. consequently, a water based dpsuspension containing 3 and 1 µm poly-crystalline diamonds on md-dur and md-nap cloths, respectively, was used for fine polishing. an md-chem cloth combined with the water-based dp-suspension containing 0.25 µm and 0.1 µm diamond particles was selected for final polishing. in such a way, two surfaces for the measuring of longitudinal (fig. 1) and transverse (fig. 2) properties of the basalt fiber were prepared. this allowed to address a potential anisotropy of the ceramic matrix in the vicinity of the fibers. when cutting the specimen in the direction normal to the fibers resulted in the average root mean square roughness rq = 95.08 nm, whereas in case of cuts along the longitudinal direction an increase of the average root mean square roughness rq = 300.75 nm was observed owing to the cleavage of fibers during preparation. the measurements were carried out analogously for both longitudinal and transverse directions start142 http://dx.doi.org/10.14311/app.2017.13.0142 http://ojs.cvut.cz/ojs/index.php/app vol. 13/2017 basalt fiber reinforced ceramics (a). (b). (c). (d). figure 1. indentation steps in longitudinal direction. (a). (b). (c). figure 2. indentation steps in transverse direction. ing with scanning the composite surfaces with an inbuild scanning probe microscopy (spm) (figs. 1a and figs. 2a) to identify location of individual indents. these are seen for the fibers in figs. 1b and 2b. the indented surfaces are evident in figs. 1c and 2c. each phase was subjected to a different loading function. separation of indents ensured no interaction of individual indents. the indented surfaces were finally scanned again to check the measurements accuracy, see e.g. fig. 1d. the load-controlled quasi-static indentation was performed with the load function consisting of “loading” and “unloading” segments lasting 5 seconds each with an in-between 300 seconds segment of “holding” time. the maximum applied force of 2500 µn was used for fiber indentation and 300 µn in case of the matrix. this procedure was applied to several positions on the sample. the resulting reduced moduli, (a). (b). (c). (d). figure 3. (a) mesoscopic 2d image of basalt fiber/ceramic matrix plain weave textile composite, (b) binary image of the 8-layer laminate cross-section, (c) description of two-layer sepuc on the mesoscale, (d) example of a real structure of yarn. corresponding to longitudinal and transverse indentations into the fibers and the matrix, were derived from unloading part of the indentation curve. next, the required young moduli were calculated based on the contact mechanics of colliding solid bodies [1, 2] assuming the poisson ratio ν = 0.24 for all phases and a diamond tip with material parameters e = 1140 gpa and ν = 0.07. taking into account all indents gave the average values of the young modulus equal to 69.68 gpa and 64.82 gpa for the longitudinal and transverse fiber directions, respectively, and equal to 58.83 gpa for the matrix. further details including basic statistics are available in fig. 1. it is not surprising that the fiber values are close confirming an essentially amorphous structure of the basalt fibers. a slight difference can be attributed to the quality of sample of the examined textile composite. an expected isotropy of the amorphous ceramic matrix has also been confirmed, since no principal difference has been observed when indenting the matrix in two perpendicular directions. application of the isotropic theory of nanoindentation to both phases is therefore adequate. 3. formulation of geometrical model during construction of the geometrical model for a computational analysis of textile composite we need to distinguish several levels corresponding to its structure. fig. 3 identifies the mesoscale structure of individual yarns and a random distribution of fibers within the yarn cross-section at the microscale level. both scales require formulation of a suitable computational model. for illustration, a cross-section of a two-ply computational model to be used on mesoscale is plotted in fig. 3c, see e.g. [3] for more details. 143 s. valentová, v. hrbek, m. šejnoha acta polytechnica ctu proceedings position red. mod. young’s mod. hardness contact depther [gpa] e [gpa] h [gpa] hc [nm] basalt fiber – longitudinal direction p127_2 67.81 ± 8.09 67.72 ± 7.68 5.27 ± 0.56 191.92 ± 6.55 p128_1 75.95 ± 6.55 75.66 ± 6.21 5.95 ± 0.88 106.09 ± 8.91 p128_2 76.85 ± 7.75 76.63 ± 7.35 5.71 ± 0.76 108.33 ± 7.98 p130_1 66.17 ± 10.44 66.08 ± 9.93 5.49 ± 0.75 161.55 ± 12.22 p130_2 71.76 ± 7.30 71.15 ± 6.92 5.79 ± 0.38 155.99 ± 5.62 p130_3 61.12 ± 6.68 60.85 ± 6.33 5.28 ± 0.74 239.19 ± 18.71 average 69.94 ± 7.80 69.68 ± 7.40 5.58 ± 0.68 160.51 ± 10.00 basalt fiber – transverse direction p120_1 74.09 ± 6.22 73.65 ± 6.22 5.88 ± 1.17 107.60 ± 11.80 p121_2 75.98 ± 3.79 75.69 ± 3.58 5.94 ± 0.54 105.58 ± 5.57 p121_3 57.31 ± 3.58 56.85 ± 3.38 5.00 ± 0.55 116.33 ± 7.25 p122_1 55.34 ± 10.55 54.80 ± 10.03 4.45 ± 1.16 126.64 ± 17.08 p122_2 68.77 ± 8.09 68.65 ± 7.68 5.15 ± 0.54 114.39 ± 6.80 p124_1 59.61 ± 11.70 59.26 ± 9.44 5.29 ± 1.22 117.95 ± 15.13 average 65.18 ± 7.32 64.82 ± 7.03 5.29 ± 1.19 114.52 ± 14.15 pyrolyzed polysiloxane matrix p127_2 67.57 ± 3.82 67.47 ± 3.61 7.29 ± 0.96 56.64 ± 4.25 p128_1 56.45 ± 3.77 55.95 ± 3.56 5.44 ± 0.44 110.65 ± 4.90 p128_2 60.20 ± 5.09 59.88 ± 4.82 6.01 ± 0.54 104.88 ± 5.21 p130_1 56.70 ± 8.53 56.22 ± 8.10 5.63 ± 1.02 109.81 ± 10.39 p130_2 58.39 ± 7.29 57.98 ± 6.91 5.68 ± 0.99 109.45 ± 11.72 p130_3 55.82 ± 5.00 55.30 ± 4.73 6.23 ± 1.08 103.85 ± 10.21 average 59.19 ± 5.58 58.83 ± 5.29 6.05 ± 0.84 99.21 ± 7.78 table 1. indentation results of individual material phases. (a). (b). figure 4. microscale level of yarn, (a) two-fiber sepuc-1, (b) four-fiber sepuc-2. in this paper the main attention is paid to the microscale, meaning the level of yarns. fig. 4 plots two types of the idealized unit cells being representative to the actual yarn. these are derived by comparing material statistics of real, fig. 3d, and artificial microstructures. such representative volume elements (rve) are call the statistically equivalent periodic unit cell (sepuc), see e.g. [4] for more details. apart from correctly capturing the real microstructure attention must also be accorded to the application of macroscopically uniform state of stress or strain and proper boundary conditions. to this end, the interested reader is referred to [3]. 4. evaluation of effective elastic properties when limiting our attention to linear elasticity the classical micromechanical models such as the moritanaka (mt) method are often preferred to more complex finite element based simulations employing sepuc. since the latter approach is adopted in the next section addressing the loading conditions beyond the elastic limits of material phases we compare both methods to support adequacy of the selected rves. the mori-tanaka averaging scheme builds upon the knowledge of the shape and orientation of the reinforcements, their material properties and volume fractions. the microstructural details are therefore not fully reflected. the material properties adopted in the presented study for individual phases are listed in tab. 2. these are derived from experimental measurements, recall tab. 1, and partly taken from literature [5]. the homogenized properties of the basalt yarn are then summarized in tab. 3. it is clear that in case of elasticity both analytical and numerical homogenization provide the results which are reasonably close. this also partly supports application of the two computational models for the nonlinear analysis, not quite possible with the mt method. 144 vol. 13/2017 basalt fiber reinforced ceramics material e g ν gf st[gpa] [gpa] [–] [n/mm] [gpa] basalt fiber – longitudinal direction 69.68 28.10 0.24 2 basalt fiber – transverse direction 64.82 26.14 0.24 0.5 pyrolyzed polysiloxane matrix 58.83 23.72 0.24 0.001 0.08 table 2. considered parameters of individual materials. model type e11 e22 e33 g23 g31 g12 cf[gpa] [gpa] [gpa] [gpa] [gpa] [gpa] [–] sepuc-1 fem 62.96 62.97 65.86 26.46 26.48 23.5 0.62 sepuc-2 fem 62.98 62.96 65.86 26.48 26.46 23.5 0.62 mori-tanaka method 62.97 62.97 65.88 26.47 26.47 23.5 0.62 table 3. effective elastic moduli. 5. damage modeling at the level of yarns henceforth, we concentrate on the description of damage evolution within a single yarn of the present textile composite. similar to the previous section, the yarn is being considered as a unidirectional composite (ud) subjected to macroscopically uniform strain. attention is accorded to transverse direction only. an improved damage evolution model proposed by fang et al. [6, 7] is applied to receive a more reliable response after the onset of cracking in comparison to, e.g. the direct stiffness reduction method. given the orthotropic character of ud composites we associate with each principal direction an independent damage parameter ωi, i=1,2,3. additionally, six damage parameters d1t , d1c, d2t , d2c, d3t and d3c are specified to further distinguish the failure in tension (t) and compression (c). their evolution is assumed in terms of an equivalent displacement χeq. the damage parameter d reflecting a material degradation with the evolution of cracks is expressed in the form dj = χj feq (χjeq − χj ieq ) χjeq (χ j f eq − χj ieq ) , j = 1t, 2t, 3t. (1) in eq. (1), χj ieq represents the equivalent displacement at the onset of failure whereas χj feq is the full damage equivalent displacement, see [6, 7] for more details. for both phases, fibers and matrix, the crack initialization is controlled by the simple maximum stress criterion. the corresponding tensile strengths are stored in tab. 2. note that compressive failure was not addressed in this study. 5.1. numerical experiment the two types of sepuc in fig. 4 representing the yarn were analyzed to show the influence of the damage parameter evolution on the mesoscopic stressstrain curve. the tensile load was controlled by the increments of the macrostrain in the two main directions perpendicular to the direction of fibers. the first figure 5. stress-strain curves under the macroscopic tensile strain in the first transverse direction. main direction corresponds to the x-axis, the second to y-axis, recall fig. 4. the resulting stress-strain curves are plotted in figs. 5 and 7. the blown-up parts in fig. 5 and 7 identify 2 points for both the sepuc-1 (points 1,2) and sepuc-2 (points 3,4) for which the evolution of damage parameter, ranging in general from 0 to 1, is plotted in fig. 6 and fig. 8 for both loaded directions. 145 s. valentová, v. hrbek, m. šejnoha acta polytechnica ctu proceedings (a) . sepuc-1 point 1 and 2 (b) . sepuc-2 point 3 and 4 figure 6. damage evolution in the matrix corresponding to the marked numbered points in fig. 5. figure 7. stress-strain curves under the macroscopic tensile strain in the second transverse direction. as you can see in fig. 6 and 8 the damage initiation in the matrix loaded in tension occurs in both directions relatively soon. this corresponds to the macroscopic stress value of about 40 mpa. the damage in the matrix phase then progresses extensively before the fiber failure takes place. the mesoscopic strength parameter st can be derived as a value of stress in the stress-strain curve after which the damage parameter starts to rapidly grow. these values are indicated in fig. 5 and 7 for individual loading directions, respectively. point out that the hexagonal arrangement of fibers in sepuc-1 should give an isotropic response. a slight difference between the two directions can therefore be attributed the unstructured (nonuniform) finite element mesh. 5.2. summary the presented paper is focused on the evaluation of the effective properties of the basalt textile ceramic matrix reinforced composites at the level of yarns with a special attention paid to the determination of the macroscopic tensile strength using computational homogenization. first the homogenized elastic moduli were obtained with the help of the mori-tanaka averaging scheme and fem simulations to support the adequacy of the two selected computational models (sepuc-1 and sepuc-2). the elastic material properties of individual phases obtained from experimental measurement and partly also from literature. then the virtual experiment on these cells was carried out, the tensile macroscopic strain was applied to these cells aiming at the determination of the macroscopic tensile strength. apart from that, the mesoscopic fracture energy can also be 146 vol. 13/2017 basalt fiber reinforced ceramics (a) . sepuc-1 point 1 and 2 (b) . sepuc-2 point 3 and 4 figure 8. damage evolution in the matrix corresponding to the marked numbered points in fig. 7. determined from the homogenized stress-strain curves, recall blue lines in figs. 5 and 7. these parameters will then enter the numerical homogenization step on mesoscale adopting more complex unit cells as the one schematically plotted in fig. 4. point out, however, that the fracture properties of both the matrix and fiber were selected in the present study ad hoc and they still need to be properly measured. to that end, nanoindentation is again expected as a suitable method of attack. list of symbols rq root mean square roughness [nm] f applied force [µn] ν poisson ratio [–] er reduced modulus [gpa] e young’s modulus [gpa] h hardness [gpa] hc contact depth [nm] g shear modulus [gpa] gf fracture energy [n/mm] st critical strength in tension [gpa] eij homogenized elastic modulus [gpa] gij homogenized shear modulus [gpa] dj damage parameter [–] χeq equivalent displacement [m] σij applied stress [gpa] �ij resulting strain [–] acknowledgements the financial support provided by the sgs projects with the application registered under the no. ohk1-014/17 and ohk1-031/17 is gratefully acknowledged. references [1] h. hertz. on the contact of rigid elastic solids. chap. 6. hertz’s miscellaneous papers, london, 1896. [2] w. oliver, g. pharr. measurement of hardness and elastic modulus by instrumented indentation: advances in understanding and refinements to methodology. journal of material research 19(1):3–20, 2004. doi:10.1557/jmr.2004.0002. [3] m. sejnoha, j. zeman. micromechanics in practice. wit press, boston, 2013. [4] j. zeman, m. šejnoha. from random microstructures to representative volume elements. modelling and simulation in materials science and engineering 15(4):325–335, 2007. doi:10.1088/0965-0393/15/4/s01. [5] m. černý, p. glogar, z. sucharda. mechanical properties of basalt fiber reinforced composites prepared by partial pyrolysis of a polymer precursor. journal of composite materials 43(9):1109–1120, 2009. doi:10.1177/0021998308097732. [6] g. fang, j. liang, q. lu. investigation on the compressive properties of the three dimensional four-directional braided composites. composite structures 93(2):392–405, 2011. doi:10.1016/j.compstruct.2010.09.002. [7] g. fang, j. liang, b. wang. progressive damage and nonlinear analysis of 3d four-directional braided composites under unidirectional tension. composite structures 89(1):126–133, 2009. doi:10.1016/j.compstruct.2008.07.016. 147 http://dx.doi.org/10.1557/jmr.2004.0002 http://dx.doi.org/10.1088/0965-0393/15/4/s01 http://dx.doi.org/10.1177/0021998308097732 http://dx.doi.org/10.1016/j.compstruct.2010.09.002 http://dx.doi.org/10.1016/j.compstruct.2008.07.016 acta polytechnica ctu proceedings 13:143–148, 2017 1 introduction 2 experimental measurements of the material parameters 3 formulation of geometrical model 4 evaluation of effective elastic properties 5 damage modeling at the level of yarns 5.1 numerical experiment 5.2 summary list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0063 acta polytechnica ctu proceedings 5:63–68, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app monitoring of track sections with long-pitch corrugation jan valehrach∗, jaroslav šmíd brno university of technology, faculty of civil engineering, institute of railway structures and constructions, veveří 331/95, 60200 brno, czech republic ∗ corresponding author: valehrach@fce.vutbr.cz abstract. the focus of this paper lies on monitoring of the track sections with rail corrugations caused by wheel sliding. short waves on the running surface of the rail head on low rail are typical for this defect. long-pitch corrugation is a significant cause of vibration and noise pollution in the railway infrastructure. therefore it is very important to understand the formation and development of this imperfection of the rails. measurements were carried out in curves of small radii on several tracks in the czech republic. in addition to the measurement of the surface of the rail head a number of supplementary parameters was evaluated for each section, such as curve radius, superelevation, track gauge etc. and the speed of passing trains as well. this paper describes rail corrugation defect and its development. our early results are presented in the conclusions. keywords: long-pitch corrugation, rail, track, measurement, speed . 1. introduction during the past recent time the railway infrastructure administrators are pushed to change their approach to the maintenance of the sections affected by the long pitch corrugation. these defects were not as frequent as they are in these days and also some of them have not been considered as a serious problem for the operability of the track. there are more factors to blame, such as increases of the track speeds, higher track loads or also (predominantly in the close vicinity of the cities) higher frequency of the passenger trains. in addition, today’s customers are not only satisfied with just the travel from point a to b or with the shortest time of transfer. they also require more comfortable travelling. it is not very difficult to notice the imperfections of the track or the increased noise levels. and there is also the second aspect of the problem, which some say is even more important – that is the influence of the imperfections for the adjacent neighborhood, such as vibrations, noise and quakes, all significantly reducing the quality of life of the people living near tracks. there is no doubt about the negative effects of the long-pitch corrugation. everyone is able to recognize that the train is much louder when passing through the curve degraded by corrugations in comparison with the passage of the same curve after the replacement of the rails. the tracks with the long-pitch corrugation require more frequent and also more costly maintenance (such as rail grinding, replacement of the rails etc.). in order to fulfil interoperable function within europe the railway network has ”to ensure a safe and uninterrupted operation of trains achieving set levels of performance” [1]. accordingly, it is necessary to guarantee required quality of the railway infrastructure and to reduce maintenance cost to minimum. with the above mentioned reasons in mind we should endavour to understand the formation and development of this defect. 2. theory of formation of long-pitch corrugation the long-pitch corrugation, sometimes called also as ”short waves”, is a sinuous inequality of the running surface of the rail. distance between wave peaks is usually between 8 and 30 cm and the depth ranges between 0.1 and 1.2 mm. a typical location, where long-pitch corrugation could be found, is the lower rail in curves of radius less than 600 m, sometimes even 700 m [2, 3]. the cause of these corrugations is the difference in length that each wheel of the wheelset rolls along the curve on the low and high rail. this difference also causes the wheel slip, occurring mostly on the low rail. slip happens not only in the longitudinal direction, but also in the lateral direction. in most cases, long-pitch corrugation does not occur alone. such defects, which are accompanying the long-pitch corrugation, can be sorted to: • those supporting the development of long-pitch corrugation (such as side wear of the outer rail), • those emerging as a consequence of long-pitch corrugation (cracked clips). since the processes surrounding the long-pitch corrugation are not documented in their entirety it is often complicated to distinguish which defect is the cause and which is the consequence. therefore, the consecuting list of defects does not take into account the above mentioned division: damaged baseplates, sleepers with transverse cracks, missing parts of fastening, cracked sleeper screws, cracked clamps in rail 63 http://dx.doi.org/10.14311/app.2016.5.0063 http://ojs.cvut.cz/ojs/index.php/app jan valehrach, jaroslav šmíd acta polytechnica ctu proceedings figure 1. long-pitch corrugation in havlíčkův brod fastening, abraded ballast behind sleeper heads etc. 3. selection of locations and measurement appropriate sections for the monitoring of long-pitch corrugation, where typical signs of this defect occur, were chosen in association with the railway infrastructure administration of the czech republic (sždc). further criteria which have influenced the selection of sections originated from the identical or at least similar characteristics and features of curves, such as: • curve radius, • track cant, • traffic load and others. subsequently, an assessment of the available data for all the individual sections was conducted. the frequency of measuring was determined with consideration to previous maintenance work, rail traffic present and planned intentions of the sždc as well. the majority of sections are double tracks thus it is possible to compare two monitored locations on one track or one location on each track, for example in a measured curve. to be able to understand more thoroughly the development of the long-pitch corrugation it is also necessary to observe other factors, which influence the defect. one of these factors is the matter of variety composition of the trains (meaning the split up between passenger trains and freight trains). variety composition of the trains complicates the whole matter, since it adds more other unknown influences. that is the reason, why we try to rule this factor out if possible. it is less difficult to succeed when considering the case with one track; hence it is more likely that the track would be passed in both directions equally. this also allows us to compare two sections within one location between themselves. on the other hand, in the case of the double track there is no certainty that both tracks would be passed with the same amount of trains or even with the same rate between the passenger and freight traffic. as an example, in some cases we can distinguish the acceleration of trains in one track and deceleration or even stopping of trains in the second track (like when approaching an entry signal of railway station). that is why it is not possible to perform comparison between the sections in different tracks within one measuring location. in both cases the monitored sections are chosen by the presume of the influence on formation and development of long-pitch corrugation, by the track alignment, by the permanent way design itself (composition of the superstructure including the effects of tested design elements such as more elastic fastening to the sleepers, clips with increased fatigue limits, under sleeper pads) and by the substructure design (e.g. the influence of the bedrock, bridge structure etc.). the sections are monitored in regular intervals depending on the present state of development of defects, traffic volume, maintenance and repair works etc. the sections with recently replaced rails allow us to observe the formation and development of long-pitch corrugation in its initial stages. measurements were repeated after 3 months. in sections, where the defects are more developed, the time between the measurements was almost doubled, since the corrugations are more noticeable (deeper and significantly better visible). that allows us to measure each wave separately. on the other hand, the measurements in sections, where defects are not yet fully developed should be performed with increased frequency, so it is easier to cover the whole development process. as you can see on figure 2, the long-pitch corrugation forms very interesting shapes which are changing fluently throughout the curve. apart from the visual inspection (which confirmed the suitability of our selected locations) all of the sections are monitored and documented using the following methods: • visual inspection and photo documentation, • measurement of track geometry by the trolley krab, • measurement of wavy deformations of the rail head by the device salamander, 64 vol. 5/2016 monitoring of track sections with long-pitch corrugation figure 2. highlighted long-pitch corrugation by contaminant in havlíčkův brod • monitoring of the speed of trains and train compositions. the track geometry is measured to describe the current state of the track, which unfortunately in many cases does not correspond to the data found in the railway documentation. another reason for these measurements is the need for additional information such as track gauge, gauge widening etc. the microgeometry of the rail head, specifically wavy deformations, was measured using the device salamander. unfortunately, in some cases it was not technically possible to measure a curve as a whole, which is the reason, why only few selected sections were measured within these locations. measured data are evaluated by the device in accordance with en 13231-3 [4] in wave lengths: • 10 ÷ 30 mm, • 30 ÷ 100 mm, • 100 ÷ 300 mm, • 300 ÷ 1000 mm. this categorization is not fully appropriate for our purposes since long-pitch corrugation can appear with a wave length shorter than 100 mm [3], particularly in small radius curves. based on the theory of formation of long-pitch corrugation [6] , our own measurements and data comparison, it can be said that the two categories for 30 ÷ 100 mm and 100 ÷ 300 mm wave lengths can be merged together for the purposes of figure 3. trolley krab (track geometry) [5] and salamander (corrugations) monitoring long-pitch corrugation. in both zones, the same or very similar wave amplitudes are achieved. apart from the monitoring of the long-pitch corrugation of the rail head, other critical elements were evaluated as important and thus were observed: • compliance with technology of construction and maintenance works, • compliance with technology of works in stations, • composition of passing trains, • operation of rail vehicles with respect to the track speed. during the measurements, the influence of the difference between the track allowed speed and actual speed of the vehicles proved to be imperative. that is the reason for closer observation of the traffic with respect to the measured speed of trains. in some cases the track allowed speed, or sometimes cogitated as a designed speed, differs significantly from the actual speed of the vehicles. this phenomenon occurs more often these days, since there is an effort of the track designers to maximize the speed limits as much as possible (they tend to ”squeeze out” the most of the infrastructure) and mostly could be found on tracks with very heterogeneous traffic flow, consisting of the various ratios of the passenger and freight trains. the more freight trains pass throughout the section of the track, the bigger is usually the difference. generally, the difference between the allowed and actual speed produces a cant excess or cant deficiency. analogically, these values grow parallelly to the difference, so when the difference increases, the values increase accordingly. therefore, a weighted average of all speed measurements in a set section was calculated, when the mass of each train was taken into consideration. the masses of the trains for the purposes of the weighted averages have been obtained by two ways. the weight of passenger trains was taken from the passenger carriages charts and from the database of the train compositions. the mass of the freight trains could be found out from the mass normatives (which is the maximum allowable mass of the whole train). the problem is that the actual load on the train is not known and thus it is not possible to know the right mass of the train. therefore the masses of the freight trains were estimated with only one value around 1000 tonnes. this approach brings the significant amount 65 jan valehrach, jaroslav šmíd acta polytechnica ctu proceedings figure 4. the processed record from the device salamander shows a track, where in one part the rails were replaced and in the other part the rails with long-pitch corrugation were left. of uncertainty to the result, but on the other hand, according to higher order masses of the freight trains in comparison to the passenger trains, the sensitivity of the result to the changes of the masses is quite small. then for the calculation of the cant excess and the cant deficiency the following scheme was used. assuming a formula for speed of trains v = √ d ∗ r 11, 8 the cant excess then applies e = d − 11, 8 ∗ v 2 r and analogically for the cant deficiency i = 11, 8 ∗ v 2 r − d ideally, the effects of cant excess and deficiency should be balanced, especially with respect to train masses. however, in some sections cant excess markedly dominates. that happens for example in the vicinity of stations, where the trains are accelerating and have not yet reached the track speed or where the greatest amount of the track load consists of slow moving trains (usually freight trains). in such locations this leads to the overload of the railway superstructure and consequently results in increased occurrence probability of the defects. figure 5. cant in track no. 1 and in track no. 2 in curve a. 4. outputs 4.1. measuring location havlíčkův brod one of the monitored track locations is a part of the track no. 324 havlíčkův brod – kutná hora in between the km 224.394 and km 225.150 in approach to the station havlickuv brod from the direction kutna hora. the location consists of a double track with consecuting reverse compound curves. first curve is for the purposes of this work named a and is closer to the station. the second curve is named b and is located further from the station in direction kutná hora. the cant of the track was measured, using the device krab, which showed a different value of the cant for each track. the results of the train speed measurements were the differing speeds of trains in both tracks and both curves. after the calculation of the weighted averages, the data revealed the presence of cant excess in both tracks and both curves. the train speed in the curve a in the track no. 1 is affected by braking/accelerating near havlíčkův brod station. some freight trains heading from kutná hora stopped at the station head in such a way that the end of the train was standing in the curve (e = 2 mm). unfortunately, this situation is a common practice in havlickuv brod. it facilitates the document handover by the stopping of the locomotive just in front of the station building. the data proved that both rails are nearly under the same load, since the weighted average e/i was close to zero. the record of the wavy deformations for wave lengths 100 ÷ 300 mm is shown on figure 6 on the first line. for the curve a on track no. 2 the weighted average equals e = 27 mm. when comparing with the track no. 1 the wavy deformations reach the amplitudes twice as high (figure 6, second line). on the contrary the cant excess markedly dominates in the curve b in comparison to the curve a. the weighted average the curve b equals e = 57 mm. the state of microgeometry as shown on figure 6 (fourth line) is practically identical to the curve a of this track. the track no. 1 of the curve b is passed with almost all trains with an cant excess (e = 72 mm) since they have to reduce the speed according to a change of allowed speed between the curves. in the past, the railway infrastructure administrator tried to improve the situation by reducing the allowable 66 vol. 5/2016 monitoring of track sections with long-pitch corrugation figure 6. category of wavelengths 100 ãů 300 mm havlickuv brod (from the top: curve a of the first track, curve a of the second track, curve b of the first track and curve b of the second track). speed to 55 kph, nevertheless this action had an adverse effect. instead of going at 70 kph, the trains are decelerating to 55 kph already during the passage through the curve b. additionally; the freight trains which are approaching the station and intended to pass through the first turnout in the side direction have to be slowed down just before the pass of the entry signal at 40 kph, which is located in the middle of the curve b. this forces such trains to travel with the reduced speed for more than 500 metres before the station. the designed cant in the curve b of 142 mm is from that reason unnecessarily large and corresponds to the recommended cant dn1 for speed 75 kph (recommended cant is the value of the cant prescribed by the standard csn 73 6360-1 [7]). the allowable speed in this curve is on the other hand only 70 kph and moreover the speed, which is achievable realistically is even less – just 55 kph. surprisingly, the wavy deformation of microgeometry is less developed in this section in comparison to other measured sections (as it could be seen on the figure 6 (third line)). one of the possible explanations of this phenomenon may be the usage of the under sleeper pads, which were installed in this curve and only in the track no. 1. according to [8], it is possible to install under sleeper pads to affect the behaviour of sleepers in a desired zone of the frequency spectrum. this is the way how it is possible to suppress the generation of vibrations and noise emissions in these frequencies. part of the energy of dynamic processes is emitted from vibrating sleepers as noise. for newly built track with under sleeper pads an increased level of noise was measured (in just units db). such a track shows over time slower development of imperfections of the rail heads. as a consequence, the noise levels for these tracks are lower than for comparable tracks without under sleeper pads. for this reason, not only the noise pollution, but also other parameters, such as vibration parameters, sleeper and rail movements and development of irregularities of the rail head are monitored on the test sections in the czech republic. the railway superstructure in the location havlickuv brod consists of the rails 49 e1, fastenings e40 with clamps e14 in the curve a and skl40 in the curve b, then sleepers b 91s/1 on both tracks. the difference between used fastening systems could not be considered as a cause of the change in the defects, since the replacement of the fastening system was the result of the development of the defects and should have been the resolution of the problem. the long-pitch corrugation was found on both tracks. additionally, massive cracking of clamps skl14 occurs in the curve b. it is surmised that high value of the weighted average of cant excess contributes to defects of the clamps. despite of the fact that the above mentioned data indicated that higher values of cant excess have the most adverse effect on microgeometry, the slowest development of the long-pitch corrugation occurred in the curve b in track no. 1. 4.2. measuring location hady the monitored location hady is the opposite to the location havlíčkův brod in the means of the unequal cant excess. in this location the cant deficiency plays the main role. this section could be found on the track from brno to česká třebová between km 161.685 and km 164.485. during the year 2015 the rails of the track no. 1 have been replaced and the development of the long-pitch corrugation thus could be observed and measured since the time zero i. e. from the perfect condition. the measuring location consists of two sections, first is located in front of the tunnel a (curve a), the second is located in front of the tunnel b (curve b). both monitored sections lay only in the track no. 1, but the line is double track. the usual direction of the passing trains in the track no. 1 is from the tunnel b to the tunnel a, which is the heading from česká třebová to brno. figure 7 shows the comparison of development speeds of wavy deformations (here represented as a percentage of the boundary overrun of the rms limit value, rms = root mean square) in both curves in location hady. initial analyses of the data show significant differences in development of the long-pitch corrugation in the curves with small radii. considering the common structure and geometrical parameters of the track (curve radii 283 m and 261 m, cant 123 mm in both curves, fastening w14, sleepers b 91s etc.) and the railway superstructure they do not exhibit significant differences. the curve b (in front of the tunnel b) embodies the weighted cant deficiency i = 67 mm. the curve a (in front of the 67 jan valehrach, jaroslav šmíd acta polytechnica ctu proceedings figure 7. comparison of rms values [%] (to the left the measuring point in front of the tunnel a and to the right the measuring point in front of the tunnel b in location hady. tunnel a) shows the weighted cant deficiency i = 47 mm. unfortunately, these values have to be considered as tentative, since the train speeds were affected by the day time in which the measuring took place. the majority of the recorded trains were passenger trains; the freight trains pass this location mostly at nights due to very strong suburban traffic. also it has to be pointed out, that the track configuration is similar to that in the location havlíčkův brod (namely the curve a of the track no. 1). there is a turnout located just behind the measuring location and freight trains approach the station through the side direction of the turnout at 60 kph at maximum. passenger trains pass the turnout in straight direction at speeds up to 70 kph ( or up to 75 kph for trains using the cant deficiency 130 mm). when we take into consideration that the measured train speeds were just tentative, the calculation offers results, in which the values for cant deficiency are logically lower (for the section in front of the first tunnel). this partial conclusion shall be confirmed or refused as soon as possible. 5. conclusions various parameters were monitored on the two measuring locations. in the every curve there could be found a different combination of factors affecting the development speed of the long-pitch corrugation and in some curves a different track parameter can be the dominant factor. nevertheless, it can be declared at the present time and based on our measurements and observations, that the speed of passing trains undoubtedly plays the significant role on the development of defects. the adverse effect of the speed occurs in two forms: • the higher the value of the speed, the more significant the development of the defect, • the higher the value of the cant inequity occurs, the more significant is the defect. at the current state of the research it might be said that the higher the speed or the bigger the difference between the actual speed of the vehicle and the designed track speed, the faster and greater the development of the defects. the curve radius has to be considered as a fixed parameter at all times, since its change in order to reduce the development of the long-pitch corrugation is in most cases nearly impossible, we have to focus on another parameters. the composition of the railway superstructure and setting of the appropriate speeds of trains and also appropriate cant values seem to be promising factors. whether these parameters alone would be sufficient the following research will show. acknowledgements the presented results were obtained with the support of centre for effective and sustainable transport infrastructure (te01020168), student’s grant tender of the faculty of the civil engineering of the brno university of technology (fast-j-15-2860) and railway infrastructure administration, state organization (szdc s. o.). references [1] r. čech. interoperabilita a rozvoj ždc [online], 2015. [2016-02-11], https://www.fd.cvut.cz/personal/ tyfal/str/predmety/ikod-pr/ikod10.pdf. [2] c. esveld. modern railway track. 2nd ed. zaltbommel: mrt-productions, 2001. doi:isbn 90-800-3243-3. [3] h. funke. broušení kolejnic, 1992. [4] en 13231-3. railway applications track acceptance part 3: acceptance of reprofilling rails in track, 2012. [5] s. s. r. kžv. krab system [online]. [2015-12-26], http://kzv.cz/english/krab.php. [6] j. valehrach. skluzové vlny teorie a praxe, 2015. doi:isbn 978-80-214-5091-2. [7] čsn 73 6360-1. konstrukční a geometrické uspořádání koleje železničních drah a její prostorová poloha část 1: projektování, 2008. [8] plášek o., hruzíková m. přínos podpražcových podložek ke zvyšování kvality geometrických parametrů koleje, in zborník prednášok z konferencie s mezinárodnou úcastou pozemné komunikácie a dráhy, environmentálna kapacita pozemných komunikacií a dráh. košice, technická univerzita v košiciach, 2006. isbn 80-8073-602-2. 68 https://www.fd.cvut.cz/personal/tyfal/str/predmety/ikod-pr/ikod10.pdf https://www.fd.cvut.cz/personal/tyfal/str/predmety/ikod-pr/ikod10.pdf http://dx.doi.org/{isbn} 90-800-3243-3 http://kzv.cz/english/krab.php http://dx.doi.org/{isbn} 978-80-214-5091-2 acta polytechnica ctu proceedings 5:63–68, 2016 1 introduction 2 theory of formation of long-pitch corrugation 3 selection of locations and measurement 4 outputs 4.1 measuring location havlíckuv brod 4.2 measuring location hady 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0050 acta polytechnica ctu proceedings 12:50–55, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app the relation between driver behavior and intelligent transport system alica kalašováa, ∗, simona kubíkováb, ján kapustab a univerzitná 1,âă010 26, žilina, slovak republic, supervisor of department of road and urban transport b univerzitná 1,âă010 26, žilina, slovak republic, phd. student ∗ corresponding author: alica.kalasova@fpedas.uniza.sk abstract. the main objective of slovakia’s transport policy is to reduce the number of traffic accidents and increase safety on our roads. implementation of intelligent transport systems presents one of the possibilities how to meet this goal. acceptance of these systems by motor vehicle drivers and other road traffic participants is necessary in order for them to fulfill their purpose. only if the drivers will accept intelligent transport systems, it is possible to flexibly and effectively manage road traffic flexibly and effectively. from the perspective of a driver it concerns, in particular, the possibility of using alternative routes when traffic accidents or other obstacles occurs on the route that would significantly affect the continuity and safety of road traffic. thanks to these technologies, it is possible to choose the appropriate route while driving, of course based on the criterion, which the driver considers the most important during the transport from origin to destination (driving time, distance from origin to destination, fuel consumption, quality of infrastructure). information isare provided to the driver through variable message signs or directly in the vehicle (rds-tmc). another advantage of intelligent transport systems is a positive impact on psychological well-being of the driver while driving. additional information about the possible obstacles, weather conditions and dangerous situations that occur on the roads as well as alternative routes are provided to the driver well in advance. this paper is mainly focused on how the drivers perceive the influence of intelligent transport systems in žilina region. keywords: safety, accidents, its. 1. introduction the slovak republic has implemented into its legislation aâălegally binding act of the european union, directive 2010/40/eu of the european parliament and of the council of 7 july 2010 on the framework for the deployment of intelligent transport systems in the field of road transport and for interfaces with other modes of transport. thereby, the slovak republic committed to meet the conditions of implementation and use of intelligent transport systems. however, effort exists to improve this situation through the adoption of act no. 317/2012 coll. on intelligent transport systems in road transport and amending and supplementing certain acts, as well as by adopting national road safety plan of the slovak republic in the years 2011-2020 and program to support the development of intelligent transport systems and national system of traffic information. the basis of these documents is a promising creation of a so-called vision of zero accident rate in the field of transport safety. in this context that regarding the european area of road traffic safety, the european commission enforces reduction of the number of road traffic fatalities by half in comparison with 2010 as a priority for years 2011 to 2020[1]. in recent years, the slovak republic managed to lower the number of traffic accidents. the ambitious objective of slovakia’s commitment under the european road safety charter to reduce the number of fatalities in road accidents to half by 2010, in comparison with 2002, when 610 people were killed on our roads, was not fulfilled. we managed to lower the number to 345 fatalities, which represents a 43% decrease. even though we failed to meet the commitment mentioned above, subsequent statistical indicators of accident rates in the slovak republic, which show a downward trend, are positively perceived in slovakia as well as in the european union[2]. figure 1 shows the decrease of accident rates on our roads from 2000 to 2015. this decrease is caused by the development and evolution of automotive technology (especially the development of active and passive safety of vehicles) as well as by better and more suitable roads, better traffic signs, proper traffic management and organizational measures and also by the use of intelligent transport systems. the rapid decline of accident rate between 2008 and 2009 was caused by the adoption of new laws and decree in our legislative related to traffic safety[3]. this paper is focused mainly on the analysis of traffic situation in žilina region, specifically on the section between the city of bytča and ivachnová village. the intensity of vehicles on this section is one of the highest in the region. the application of intelligent transport systems is one of the most effective 50 http://dx.doi.org/10.14311/app.2017.12.0050 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 the relation between driver behavior and intelligent transport system figure 1. evolution of accident rates in the slovak republic from 2000 to 2015. city road name intensity traffic accidents ružomberok d1 19636 3 hlboké nad váhom-strečno i/18 26087 86 šútovo-kraľovany i/18 21225 5 kraľovany-dolný kubín i/70 13766 8 vrútky-kraľovany i/18 21758 33 kraľovany-sliače i/18 21861 31 žilina-terchová ii/583 5366 39 terchová-párnica ii/583 3843 5 table 1. traffic accident rates on the selected sections of road network in 2014. ways how to improve traffic situation on this road section. 2. characteristic of the selected section of road network selected section of road network consists of the first and second class roads and d1 motorway. the statistics of traffic accidents based on the section and type of road are listed in the following table 1. presented data are from 2014. the table shows that the most traffic accidents in 2014 on a selected road network were documented on the section from hlboké nad váhom to strečno (road i/18), where the highest intensity of vehicles per 24 hours was also recorded. 2.1. measurement of road network safety indicators of accident rates usually determine the safety on roads. safety is the result of risk and level of risk influence. the risk is expressed by the number of traffic accidents and the level of risk influence is specified through intensity. the indicators always refer to the same period of time, which is in this case one year. this enables us to determine the frequency of traffic accidents. f requency = n/t (1) where: n – number of traffic accidents, t – time period [year]. relative accident rate represents the probability of occurrence of traffic accident on individual sections. it is calculated as the ratio between the number of traffic accidents and traffic performance. when calculating this indicator, it is necessary to distinguish whether it is a calculation of relative accident rate at intersections or on sections between intersections[4]. r = n 365 ∗ rp di ∗ t ∗ 106 (2) for intersections r = n 365 ∗ rp di ∗ l ∗ t ∗ 106 (3) for sections where: n – number of traffic accidents, rp di – annual average of daily intensities, t – time period, l – length of examined section. another indicator that is used to assess safety is the density of traffic accidents. it refers to the number of road accidents and to the section length, thus intensity is not taken into account. it is calculated according to the formula: h = n l ∗ t (4) where: h – density of traffic accidents, n – number of traffic accidents, t – time period, l – length of examined section. frequency of accidents, relative accident rate and density of accidents were the indicators used in this paper. these indicators were calculated based on the statistical data from 2014 and they are listed in table 2[5, 6]. as shown in table 2, relative accident rate which depends on the transport performance is not critical in any case. however, it should be noted that property damage only (pdo) accidents were not taken into account since these statistics were not available. otherwise, the values of relative accident rate would be much higher. the highest value of density of traffic accidents (3,1 accidents per one kilometer) was recorded on the stretch from hlboké nad váhom to strečno, which is the section of road i/18. also in this case, property damage only accidents were not included in the calculation. this analysis shows that it is necessary to solve the traffic situation on the given section mainly through long-term organizational measures as well as by the application of intelligent transport systems. voluntary acceptance of intelligent transport systems by drivers and other road traffic participants is necessary in order for it to be effective. a questionnaire survey, whose main goal was to identify the requirements and needs of road users as well as the level of perception of intelligent transport systems was conducted within the žilina region. 51 a. kalašová, s. kubíková, j. kapusta acta polytechnica ctu proceedings section rpdi [veh/24] l [km] n [number of accidents] r [%] h [number of accidents/km/year] ružomberok 19636 3,06 3 13,67899 0,980392157 hlboké nad váhomstrečno 26087 27,7 86 32,60631 3,104693141 šútovo-kraľovany 21225 5,2 5 12,41155 0,961538462 kraľovany-dolný kubín 13766 18,1 8 8,79652 0,44198895 vrútky-kraľovany 21758 15,7 33 26,46686 2,101910828 kraľovany-sliače 21861 27,2 31 14,28334 1,139705882 žilina-terchová 5366 24,5 39 81,27463 1,591836735 terchová-párnica 3843 20,4 5 17,47337 0,245098039 table 2. accident indicators. 3. analysis of perception of intelligent transport systems from the perspective of drivers the survey was conducted electronically through a website, as well as by personal conversations with residents who possess a valid driving license. data was collected during one month, from april 13, 2015 to may 19, 2015. subsequently, the data were was analyzed and evaluated graphically and in tables. the tool called ,"sample size calculator" was used to determine the minimum sample size that is required for sufficient explanatory power of the survey, in which we set up the population of registered drivers in žilina region to 414âă678, we chose 95% confidence interval and permissible margin of error was 5% (fig. 2). the minimum sample was set to 384 respondents. a total of 519 respondents participated in the survey and therefore we can conclude, that the survey has sufficient explanatory power. figure 2. sample size calculator. the first part of the questionnaire was focused on the categorization of respondents according to age, sex and also according to the fact whether the respondent is an active or passive driver (fig. 3). the purpose and regularity of trips made during the week were also examined (fig. 4). when choosing the route, every driver prefers the figure 3. percentage of active and passive drivers. route that best suits his needs and criteria. essential criteria for choosing the most appropriate route are fuel consumption, time accessibility of the destination point, distance between origin and destination, the danger of traffic accidents and so on. the evaluation scale was as follows: 0 – irrelevant, 1 – the least significant, 5 – the most important. weighted arithmetic means of importance of the individual criteria are calculated from these values and they are graphically processed in table 3, which shows that the most important criterion for men in both age categories is the quality of transport infrastructure. time accessibility of destination point is the most important for women. traffic congestions and delays on roads can be eliminated through the opportunity to use an alternative route, to which the vehicle will be diverted via information systems installed directly in the vehicle (rds-tmc) or through information systems situated on roads that contain information for the entire traffic flow (variable message signs). in the questionnaire survey, respondents were also asked whether they would be willing to take an alternative route offered by the information system in case of traffic accident or the occurrence of other obstacles on their original route. the results are shown in fig. 5. the most frequent answer of the respondents was that they are willing to accept an alternative route even if the distance to the destination would rise by 52 vol. 12/2017 the relation between driver behavior and intelligent transport system figure 4. purpose of transport and routing of traffic flows. criteria men 18-26 men 27-65 women 18-26 women 27-65 fuel consumption 3,59 3,52 3,10 3,75 travel time 3,79 3,95 3,96 4,13 distance between origin and destination 3,27 3,39 3,40 3,75 quality of road infrastructure 3,80 3,97 3,90 4,13 possibility of congestion 3,48 3,66 3,42 3,50 possibility of road accident 3,23 2,71 3,52 3,69 table 3. criteria for route selection – weighted arithmetic mean. figure 5. acceptability of alternative route due to the length of detour (left) and with respect to the prolongation of travel time (right). 53 a. kalašová, s. kubíková, j. kapusta acta polytechnica ctu proceedings more than 2 km and time accessibility would increase in the interval from 6 to 10 minutes. sufficient awareness of drivers concerning road traffic situation is another way how to avoid roads with high intensities, congestions or road sections where a traffic accident occurred. therefore, when planning the transport, it is appropriate to have the actual current traffic information, which can be detected by the driver before or during the trip through information technologies. according to the survey, the most frequently used means of obtaining traffic information is the radio, as can be seen in figure 6. figure 6. information systems used and their share in obtaining traffic information. in the next question, we asked whether the drivers would appreciate intelligent transport systems, which would alert them about the possible dangers and which of the dangerous situations are the most important to them. the survey shows that women consider the information about black ice and fog to be the most important and the least important are for them warnings concerning dead-end streets. men consider the information about traffic accident and black ice to be the most important and the least important are for them warnings about dead-end streets and one-way roads. the evaluation scale was as follows: 0 – irrelevant, 1 – the least significant, 5 – the most important. weighted arithmetic means, which is processed graphically in figure 7, were calculated based on the values obtained. figure 7. evaluation of the importance of warning for driver about dangerous traffic situation. the last part of the questionnaire was devoted to active safety components. it was examined, whether the drivers consider the presence of selected components of active safety to be necessary. active safety components are defined as systems and devices that help to prevent the occurrence of traffic accidents and thus increase the safety of road transport. table 4 shows which systems the respondents consider to be the most important. the evaluation scale was as follows: 0 – irrelevant, 1 – the least significant, 5 – the most important. the weights of importance were evaluated based on the weighted arithmetic mean. an overall analysis of the questionnaire shows that the drivers are open to new possibilities of using intelligent transport systems, especially with regards to the increase of their own safety while travelling by a motor vehicle. 4. conclusion building telematics technologies is expensive. it is therefore not possible to apply them to all roads. however, they should form an essential part of motorways, expressways and first class roads, mainly because of high traffic intensity on these roads. several cities, where it will be necessary to solve growing traffic problems by building intelligent transport systems, were chosen, based on the analysis of traffic – transport processes, which was performed within the implementation of the national system of traffic information in individual regional capitals and in some larger cities. sections of road network that pass through the cities of žilina, martin and ružomberok were examined in this paper. these sections have the highest vehicle intensities and it is therefore necessary to solve the traffic issues through the use of intelligent transport systems. telematics technologies have a positive impact on traffic and its characteristics (intensity, density, speed, etc.) reduction in the number of traffic accidents can be achieved by reducing the intensity on congested road sections. the use of variable message signs to inform drivers about the current traffic situation and possible obstacles that affect the road capacity appears to be a suitable tool for reducing the risk of traffic accidents. intelligent transport systems have a positive effect on psychological well-being of the driver while driving. awareness plays an important role not only in warning about current traffic situation but also in planning the transport before it actually starts. drivers can check road traffic situation thanks to the topical and real-time information. road safety may be improved by reducing the number of traffic accidents and minimize their consequences and thus society can save hundreds of thousands of euros. the main benefit of reducing the number of traffic accidents remains minimization or complete elimination of the negative impacts of traffic accidents on mental and physical aspects of people involved in traffic accidents as well as their relatives. 54 vol. 12/2017 the relation between driver behavior and intelligent transport system ecall 3,54 3,55 4,13 3,94 parking assistant 1,89 2,26 2,63 3,00 cruise control 2,66 2,76 2,85 3,06 stop & go 1,66 1,61 2,67 2,94 automatic system for the detection of congestions 2,86 3,11 3,35 3,88 lane departure 2,56 3,16 3,63 3,81 maintaining safety distance 3,86 2,97 3,23 3,56 traffic signs 2,42 2,37 2,60 3,19 table 4. evaluation of the importance of active safety systems’ application in vehicles. references [1] a. kalašová, j. paľo. dopravné inžinierstvo – organizácia a riadenie. žilina: edis vydavateľstvo žu v žiline, 2003. 162 s., isbn 80-8070-076-1. [2] národný plán sr pre becep 2011-2020, july 2014. [online] available from: http://www.becep.sk/file_download/6/narodny_ plan_sr_pre_becep_2011_2020.pdf. [3] j. ondruš, s. kubíková. interakcia bezpečnosti a inteligentných dopravných systémov. vedecký – recenzovaný online časopis – svet dopravy, july 2014. issn 1338-9629, available from: http: //www.svetdopravy.sk/interakcia-bezpecnosti-a\ -inteligentnych-dopravnych-systemov. [4] centrum dopravního výzkumu. jak lze měřit bezpečnost? [online], available from: http://www.czrso.cz/clanky/ jak-lze-merit-bezpecnost-1-cast. [5] road accident investigation guidelines for road engineers, 2007. available from: http: //www.who.int/roadsafety/news/piarc_manual.pdf. [6] m. sørensen, r. elvik. black spot management and safety analysis of road networks – best practice guidelines and implementation steps. tøi report 919/2007, oslo, july 2014. isbn 978-82-480-0809-5, available from: http: //www.toi.no/getfile.php/publikasjoner/t%d8i% 20rapporter/2007/919-2007/919-2007-nett.pdf. 55 http://www.becep.sk/file_download/6/narodny_plan_sr_pre_becep_2011_2020.pdf http://www.becep.sk/file_download/6/narodny_plan_sr_pre_becep_2011_2020.pdf http://www.svetdopravy.sk/interakcia-bezpecnosti-a\-inteligentnych-dopravnych-systemov http://www.svetdopravy.sk/interakcia-bezpecnosti-a\-inteligentnych-dopravnych-systemov http://www.svetdopravy.sk/interakcia-bezpecnosti-a\-inteligentnych-dopravnych-systemov http://www.czrso.cz/clanky/jak-lze-merit-bezpecnost-1-cast http://www.czrso.cz/clanky/jak-lze-merit-bezpecnost-1-cast http://www.who.int/roadsafety/news/piarc_manual.pdf http://www.who.int/roadsafety/news/piarc_manual.pdf http://www.toi.no/getfile.php/publikasjoner/t%d8i%20rapporter/2007/919-2007/919-2007-nett.pdf http://www.toi.no/getfile.php/publikasjoner/t%d8i%20rapporter/2007/919-2007/919-2007-nett.pdf http://www.toi.no/getfile.php/publikasjoner/t%d8i%20rapporter/2007/919-2007/919-2007-nett.pdf acta polytechnica ctu proceedings 12:50–55, 2017 1 introduction 2 characteristic of the selected section of road network 2.1 measurement of road network safety 3 analysis of perception of intelligent transport systems from the perspective of drivers 4 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0016 acta polytechnica ctu proceedings 11:16–21, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app railtopomodel and railml 3 in overall context adam hlubuček ctu in prague, faculty of transportation sciences, konviktská 20, prague, czech republic correspondence: hlubuada@fd.cvut.cz abstract. this paper aims to provide a brief insight into the uic railtopomodel and railml initiative. concerning the railml exchange format, mainly the current form of forthcoming third edition, involving especially the infrastructure schema, based on the railtopomodel modelling principles, is taken into account. when rewarding, the comparison between railml 3 and the previous railml 2.3 version is given. at the end, the author focuses on selected issues of possible applicability of these tools for the needs of such systems as the automatic train control, automatic train operation, prediction of train etc. if essential, some extensions of the structures are proposed. keywords: railtopomodel, railml, railway infrastructure, information systems. 1. introduction railml [1] is an open source exchange format used in the railway branch to ensure a data transfer among different software applications. the railway markup langue, based on xml, was originally designed to enable data exchange in the scope of timetabling. later on, the specifications [2] were extended to the subsystems of rolling stocks and infrastructure. in these days, also the schema of interlocking is being developed in relation to oncoming railml 3 release. another milestone, as regards the railml 3, is a significant reedition of the infrastructure concept. whereas the principles of modelling of the railway topology were rather casually defined within the previous versions, railml 3 is supposed to be strongly founded on the uic railtopomodel [3] which provides the rules how to describe the infrastructure in predefined way, nevertheless, allowing all kinds of user-specific extensions. the railtopomodel initiative was introduced in 2013 [4] in order to create a common standard in railway infrastructure modelling. the generic railway topology model aims to meet multiple needs in the field of railway data exchange, ensured by the means of the railml 3 format. in 2016, the railtopomodel was released in the form of uic international railway standard 30100 [5]. 2. the railtopomodel overview the uic railtopomodel [3] is a topological model of railway infrastructure, based on a „connexity graph“. describing a railway topology, two basic types of elements can be taken into account: the linear elements (e.g. tracks and lines) and nonlinear elements (e.g. operational points). these elements allow describing the infrastructure at different levels of detail, in order to be suitable for different specific purposes. regarding the “connexity graph”, the elements of both these categories are supposed to be modelled as the nodes, whereas the edges in the node-edge diagram represent the physical connections between these particular interconnected net elements. the linear elements shall have their orientation defined. therefore, when concerning the (again orientated) relation between two of them, it should be in both cases apparent, whether the beginning or the end of the net element is taken into account. this generic principle is illustrated in the figure 1 considering the tracks to be the nodes. in order to express the fact, whether a train is able to pass from one net element to another, additional information is necessary to be given. for this reason, the relations between net elements shall be described also by the navigability attribute, besides, concerning the direction of allowable train movement. [5] in this way, a feasible passage through a switch from one track to another can be modelled as well as the line change when passing through an operation control point. currently, the model enables to express the topology including positioning and location of net entities to be projected to the topological layer taking into account different views. there are several predefined levels of detail (macro, meso, micro) at which a railway network can be described. a user may utilize (or start with) any of them, depending to his specific use case. furthermore, he is authorised to create own levels, to be appropriate for his needs (e.g. nano or corridor level). these levels of detail should be vertically interconnectable by the means of aggregation and disaggregation principles. therefore, when passing from one level to another, sometimes an extra transition level is needed. except the different views of the topology, the particular level enables to deal with particular entities in different ways. for instance, it allows displaying a certain entity as a section at a detailed level, as a spot at a less detailed level and not to include it to the least detailed level at all. [5–7] 16 http://dx.doi.org/10.14311/app.2017.11.0016 http://ojs.cvut.cz/ojs/index.php/app vol. 11/2017 railtopomodel and railml 3 in overall context figure 1. an application of the railtopomodel principles on an example of a railway station with five simple turnouts and one single slip modelled by the means of positioned net relations. later on, the model together with railml 3 should be extended so as to enable digital continuity for infrastructure life cycle and operation supporting bim development and using railml 3 for web services as well. it should be able to provide a unique continuous network description, not only taking into account past, current and future states, but also involving the project management aspect of the lifecycle dimension. [6] concerning the objects being modelled, expressing their validities (active / non active), variants (alternative current states) and versions (states evolving over time) should be possible. [4] 3. railtopomodel packages 3.1. base currently, the railtopomodel consists of four packages. concerning the uml class diagram designed to express its structure, there is a base network domain consisting of the base object, network, level network and network resource data classes, defining the general overview of the network being modelled, including the modelling principles used. the classes of the other subsystems are mostly some kind of extensions of the network resource class, eventually of the base object class itself. the base object class specifies the fundamental attributes to be common for all the derived xml elements (in order to express its unique id, name and validity). as regards the id, especially the uuid should be used. [5, 8] 3.2. topology the topology package is especially based on the net elements and relations which are modelled by the means of node-edge diagram. the aggregation principle (of net elements) is managed within the topology domain as well. particular net elements can be grouped to larger units, creating new net elements, which are supposed to be used to express the topology at less detailed level. for instance, several tracks (linear net elements in the scope of micro-topology) can be aggregated in order to form a station area, further regarded to be an operational point (a nonlinear element in the scope of macro-topology). nevertheless, the railtopomodel defines the data classes using the most general approach (not yet specifying the particular elements of infrastructure). [5, 8] 3.3. positioning the positioning package enables to define several positioning systems (linear as well as geometric – both the geographical and these ones used to display particular entities on a monitor) as well as the corresponding coordinates. to a particular net element, several of these positioning systems can be associated, nevertheless every positioning net element involves own intrinsic positioning system, whose coordinates takes values in the range of <0;1> referring to a relative position within the element. [5, 8] in order to be able to express the intrinsic position in absolute terms, the attribute of length related to the oriented net element would have to be specified in addition. 3.4. net entities the last subsystem deals with net entities. the net entities represent these object and characteristics of the infrastructure, which can be connected to the undelaying topology layer using positioning systems. differently from the previous railml versions, enabling only to link these object to a one certain track forming its data substructure, the railtopomodel (as designed for railml 3) allows them to be handled separately without relation to the topology (for instance to describe an asset, which has not been fitted to the infrastructure yet). closer specification of these net entities is provided by the railml 3 infrastructure schema. together with the entities, their location in relation to net elements and positioning systems is managed. an entity location can be a spot location, linear location or area location, while an entity is allowed to have several locations (including both different and the same types). [5, 8] 17 adam hlubuček acta polytechnica ctu proceedings figure 2. illustration of different approaches in locating of entities in the scope of railml 2 and railml 3. 4. railml evolving over time currently, railml 3 is available for testing purposes only. in 2016, the railml 3.0.is4 alpha version was released containing xsd files (defining allowable structure of particular xml documents) especially regarding the railtopomodel schema (introduced with the railml 3 release) and the infrastructure schema. [8] the railml 3.0 releases are designed for possible railml 3 users, who are asked to compare it to their needs and to formulate a “use case” defining what kind of data they intend to exchange. these “use cases” [9] serve as the basis for further development of the railml 3 infrastructure schema. by means of that approach, the creators want to avoid accumulation of such data items which no one needs. the railml 3 infrastructure schema, as regards the 3.0.is4 release, consists of following subsections: the topology, the coordinates, the operation, the geometry, the assets and the infrastructure parameters (while the last one was removed with the 3.0.5 version). [8] 4.1. topology concerning the previous railml versions, [10] topology information is stored within the particular track elements using mutual connectors situated at the track ends and track begins or within switch points. the points defining a switch are allowed to be situated both at the track begin and at a certain distance from the track begin. such approach makes the model ambiguous, as it is not clear how to segment the tracks, not to mention confusing manner in which railml 2 describes the branching tracks of a switch. [7, 10] starting from the third version, the topology is considered separately and based on the abovementioned principles of railtopomodel. concerning the micro level, a connection between tracks is allowed to be situated at the track begin or at the track end only, which admits the values (expressed in relative units) of either 0 or 1. [8] among others, lengths of tracks are not necessary to be given in order to describe “the pure topology” properly. the same approach may be used for any other level of detail. nevertheless, there are still some questions under discussion within the railml community. railml 3 should support precious description of the track geometry. it has not been decided yet, whether the geometry will be considered separately or if it should depend on the topology. [6] 4.2. positioning systems and coordintes in railml 3, the positioning systems are managed in the scope of the coordinates subsection. previous versions are only able to express position from the beginning of a track completed with information of absolute position and geographical coordinates. these are expressed as the attributes or subelements of the elements representing the object being placed. [10] the absolute position should represent traditional mileage, which indicated values are not guaranteed to correspond to real distances. it happens for historical reasons (due to reconstructions of railway lines) and it is caused by different lengths of parallel tracks in horizontal curves and in the rail branching areas. in railml 2, there is a mileage change element enabling to express mileage jumps and overlaps being a part of the track topology assigned to a particular track. within railml 3, the issue is addressed in more complex way, following the railtopomodel principles again. a user is allowed to define different linear and geometric positioning systems using line, geometric and screen coordinates. the mileage irregularities are resolved by the means of binding different linear positioning systems. regarding the intrinsic positioning system, inseparably associated to given net element, as already mentioned, the position is measured in relative values from 0 to 1). [6, 8, 10] 18 vol. 11/2017 railtopomodel and railml 3 in overall context 4.3. net entities: operational entities, geometry entities and assets the most of particular subsections of the railml 3 infrastructure schema deal with located net entities, to be fit to the topological layer. these are the operational entities, geometry entities and assets (in railml 3.0.is4 mostly representing technical facilities of railway and with 3.0.5 version renamed to functional assets in order to be distinguished from physical assets). as regards railml 3.0.is4, there are following entities included, while users can add any others (not to be a part of the common basis) to meet their needs. [6, 8] in terms of further releases, the lists are expected to be extended, based on the “use cases”. [9] • operational entities . line . operational point . border . speed profile . speed restriction . track • geometry entities . horizontal curve . gradient curve . superelevation curve . track geometry point • assets . bridge . buffer stop . derailer . isolation rail joint . level crossing . platform . platform edge . service section . signal . stop post . switch . train detection element . train protection element . tunnel for comparison, there are two major groups of entities (the track elements group and the ocs elements group) in railml 2 creating data substructures of each track element. their location within their parent track shall be given by means of the position attribute, expressing the distance from the beginning of a track in metres. in the case of linear character of the entity being placed, railml 2 uses two different approaches how to represent it; either to determine the spots where the data change (e.g. the radius changes, speed changes. . . ), or to determine the spot location of the object itself, possibly completed with the attribute of length (e.g. the platform edges, bridges, level crossings. . . ). in some cases, it is essential to determine the considered direction. [10] the railml 2 way of description, indeed, sometimes implies data inconsistency. the data could be misunderstood. the fact that all entities are strictly allocated to one specific track may cause some difficulties concerning the infrastructure description. for instance, how to define and describe track circuits when only their borders can be marked out? structural object as bridges, tunnels and level crossings often belong to more than one track. in the scope of railml 2, every affected track needs to involve an extra instance of related data classes, although the original real world object is always the same (as seen in the figure 2). [7, 10] railml 3, building on the railtopomodel principles, offers solutions of these issues. not only that it enables spot, linear and aggregated location of net entities, but it also allows them to be multiply located (see again the figure 2). locating of all the net entities is managed in the same way, no matter what kind of entity it is. [8] every xml element representing a net entity can contain unlimited amount of spot, linear and aggregated location elements involving location spots and linear sections expressing different locations of the entity. regarding the aggregated location, it can be expressed both by spots and linear sections defining an area within the railyard. the location can be also represented by point and linear types of coordinates. their definition is taken from the gml geometry schema. it means that the net entities are already not necessarily assigned to one specific track. [7, 8] the xml elements describing the entities of the same type are included in common container elements within relevant subsection. in some cases, more complex substructure is used (especially as regards the signals, nevertheless, some of their items could be possibly managed within the interlocking schema). concerning the assets, their operating time, constriction time and blocked time is possible to be expressed, in addition. prospectively, a link of these functional assets with physical assets to be a part of inventory management is intended to be ensured. [6, 8] 4.4. infrastructure paramteres the last category, also enabling to describe characteristics of railway infrastructure, used in railml 3.0.is4 is the subsection of infrastructure parameters. the following (again extendable) list of the parameter classes is given, as regards the 3.0.is4 release. [8] • infrastructure parametres . track gauge . speed electrification . weight limit . train protection . clearance gauge . train radio . track alignment 19 adam hlubuček acta polytechnica ctu proceedings in railml 2, the most of these parameters can be expressed in two different ways: through infrastructure attribute groups (to be assigned to a certain track) or as track elements (changes within a track). [10] the railml 3.0.is4 approach is closer to the first attitude mentioned. the infrastructure parameters are managed in rather different way than the net entities. every net element (as be a part of the topologic layer) can refer to unbounded number of infrastructure parameters representing specific track conditions. besides, it means that the parameters are valid for the relevant net entity (e.g. a track or line section) as a whole. except they cannot be positioned, the related xml elements are handled in the same manner as the net entities. [8] apparently, there was an effort to remove the redundancy. unfortunately, it would mean that these global parameters cannot be changed within a particular net element. indeed, in some cases it is not essential; however, for example, as regards the electrification, it could cause serious difficulties. for that reason, based on discussions within the railml community, [11] the infrastructure parameters subsection was cancelled with railml 3.0.5 (released in march 2017) and the entities included was moved to the functional assets group (thereby becoming net entities). at the same time, the network location was introduced in order to represent the intended function of the cancelled subsection. [8] 5. selected application issues the railml 3 exchange format is supposed to find application in interconnection of such it systems as infrastructure registers, maintenance systems, evidences of assets, construction projects, visualisation of railway infrastructure (georeferenced documents, gis applications, schematic track plans, longitudinal sections), simulations of railway operation, calculations of energy consumption, capacity planning, information systems for operative control, predictions of train running, automatic routing, automatic train control, automatic train operation and many others. apparently, it is big challenge for developers to meet all the multiple needs. certainly, the alpha release does not cover all the desired aspects, as surely as it will be gradually complemented reflecting particular use cases. [9] 5.1. railml for atc? as the infrastructure description used by railml is essentially based on real track lengths, not only to use absolute mileage which can be misleading, railml seems to be very useful tool to provide data for such applications that operate with actual distance travelled by a train. among other applications, it is essential for the ertms / etcs [12] as an automatic train control system using balises to be reference points for measuring the distance travelled. the balises and balise groups are included in the railml 2 specifications, nevertheless, railml 3 have not been saturated with such entities yet. perhaps, it is caused by the unresolved issue in which way to perform referencing between the balises and balise groups elements. [8, 10] except the balise locations, other additional data are needed, including locations of important objects, speed restrictions, slopes, track conditions (change of traction systems, powerless sections, non stopping areas, sound horn commands...) road suitability parameters (loading gauge, traction system, axle load category...). many of them have already been included into the railml 3 infrastructure schema, nevertheless, some of them should be added or revised. [8, 11] 5.2. transfer times planning concerning predictions of train running (and in some cases capacity planning as well), not only the railway infrastructure itself must be taken into account, but also some related facts are important. for instance, when ensuring correspondences among trains, transfer times are often critical. among others, they depend on parameters of the paths connecting the platform edges. when there are platforms with level access, the transferring passengers are influenced by passing trains (and conversely). this fact shall be taken into account when assigning the platform edges and tracks to the trains as well (for instance, not to allow a passing train run over the foot level crossing being used by passengers). not to mention the vital importance of the foot level crossing location register for such application as the automatic train operation because of the need to ensure human safety. to cope with these cases, the model and the exchange format would need to be significantly extended. 6. proposal of a rtm based passenger movement model the abovementioned issue is more complex, as the passenger routes (connecting particular platforms together and with station building and other important locations in a station) consist of different elements (e.g. platform ramps, stairways, lifts, underpasses, overpasses, foot level crossings, footpaths). all these facilities could be expressed using railml 3 as net entities (although the railml 3 alpha release, for instance, enables to express the existence of a lift on a platform by the means of its attribute only). these platform accessories and related facilities should be interconnectable and describable by such attributes expressing their usability by persons with reduced mobility and other important parameters so as to enable searching suitable routes connecting particular “passenger access nodes” (in the sense of significant spots assigned to platforms, station building and other net entities). even such situations as a lift being out of service should be taken into account. 20 vol. 11/2017 railtopomodel and railml 3 in overall context figure 3. an author’s proposal of railtopomodel application in the scope of passenger routing issue exemplified for the case of a railway station equipped with platforms with level access. there are several possible solutions how to meet these needs. one of them is to apply railtopomodel again, this time not to connect pieces of railway infrastructure but the abovementioned “passenger access nodes”, henceforth to be considered as a new sort of the net elements (see the figure 3). this approach, in general, would also enable to model emergency access roads to railway tunnels, for instance, and to cope with other related issues. 7. conclusions railtopomodel and the corresponding data exchange format railml 3 have a great opportunity to become very useful tools in describing railway infrastructure and ensuring the rail data sharing. involving also the schemas of timetable, rolling stocks and prospectively interlocking, the railml language covers almost the whole range of the railway branch. although the railway system is based on relatively similar principles across the whole europe, there are many nationally specific details that should be addressed in a coordinated manner in order to make the constantly evolving railml system universal enough to enable different software application to consistently communicate with each other. to make the system as usable as possible, the potential users should formulate their “use cases” [9] and be in contact with the railml community creating the specifications. then they should be able to influence the final form of the railml 3 first public release in order to be suitable for their needs. nevertheless, the very special matters can be solved individually by the means of user-specific extensions. within this document, some of possible extensions were outlined as well. references [1] railml. railml [online]. https://www.railml.org/. [2] railml. the railml subschemas [online]. https://www.railml.org/en/user/subschemes.html. [3] railtopomodel. railtopomodel [online]. http://www.railtopomodel.org//. [4] uic. feasibility study uic railtopomodel and data exchange format [pdf], 2013. http://www.railtopomodel.org/files/download/ railtopomodel/270913_trafit_ finalreportfeasibilitystudyrailtopomodel.pdf. [5] uic. uic international railway standard [pdf], 2016. http://www.railtopomodel.org/files/download/ railtopomodel/270913_trafit_ finalreportfeasibilitystudyrailtopomodel.pdf. [6] railml, uic. 7th railtopomodel and 30th railml conference, paris, 2016. [7] v. p. kolmorgen, c. rahmig. railml/rtm explanation session, prague, 2017. [8] railml. railml schema version 3.0.is4 [xsd], 2016. [9] railml. use cases [online]. https://www.railml.org/en/user/use-cases.html. [10] railml. railml schema version 2.3 [online], 2016. https://www.railml.org/files/download/schemas/ 2016/railml-2.3/documentation/railml.html. [11] m. karlsson, c. rahmig. [railml3|alpha] modelling of track conditions [online], 2016. http://www.railml. org/forum/index.php?t=msg&th=474&start=0&. [12] ertms. the european traffic management system [online]. http://www.ertms.net. 21 https://www.railml.org/ https://www.railml.org/en/user/subschemes.html http://www.railtopomodel.org// http://www.railtopomodel.org/files/download/railtopomodel/270913_trafit_finalreportfeasibilitystudyrailtopomodel.pdf http://www.railtopomodel.org/files/download/railtopomodel/270913_trafit_finalreportfeasibilitystudyrailtopomodel.pdf http://www.railtopomodel.org/files/download/railtopomodel/270913_trafit_finalreportfeasibilitystudyrailtopomodel.pdf http://www.railtopomodel.org/files/download/railtopomodel/270913_trafit_finalreportfeasibilitystudyrailtopomodel.pdf http://www.railtopomodel.org/files/download/railtopomodel/270913_trafit_finalreportfeasibilitystudyrailtopomodel.pdf http://www.railtopomodel.org/files/download/railtopomodel/270913_trafit_finalreportfeasibilitystudyrailtopomodel.pdf https://www.railml.org/en/user/use-cases.html https://www.railml.org/files/download/schemas/2016/railml-2.3/documentation/railml.html https://www.railml.org/files/download/schemas/2016/railml-2.3/documentation/railml.html http://www.railml.org/forum/index.php?t=msg&th=474&start=0& http://www.railml.org/forum/index.php?t=msg&th=474&start=0& http://www.ertms.net acta polytechnica ctu proceedings 11:16–21, 2017 1 introduction 2 the railtopomodel overview 3 railtopomodel packages 3.1 base 3.2 topology 3.3 positioning 3.4 net entities 4 railml evolving over time 4.1 topology 4.2 positioning systems and coordintes 4.3 net entities: operational entities, geometry entities and assets 4.4 infrastructure paramteres 5 selected application issues 5.1 railml for atc? 5.2 transfer times planning 6 proposal of a rtm based passenger movement model 7 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2018.19.0030 acta polytechnica ctu proceedings 19:30–35, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app uncertainty evaluation in criticality calculations using the tsunami methodology matěj šikla, b, ∗, františek havlůjb a department of nuclear reactors, faculty of nuclear sciences and physical engineering, czech technical university in prague, v holesovickach 2, prague 8, cz b department of reactor physics and fuel cycle support, nuclear research institute rez near prague, cz ∗ corresponding author: siklmate@fjfi.cvut.cz abstract. procedure of criticality calculations and uncertainty evaluations currently has several insufficiencies which could lead to potential non–conservativeness. paper discuss and describes possibilities of elimination of these insufficiencies and thereby ensuring safely sub–critical results of calculations. selection of experiments for validation based on similarity coefficients between systems acquired from tsunami-ip is used for this purpose. tsunami-ip operates with sensitivity of multiplication factor to cross sections of individual isotopes and reactions between compared systems. moreover, tsunami-ip determines uncertainty of calculations brought in using insufficiently similar experiments. within research, effects of variable parameters used in calculations inputs are investigated. in conclusion optimized suggested approach for validations and uncertainty evaluation using tsunami module from scale based on knowledge acquired within research is described. keywords: validation, subcriticality, tsunami, similarity between systems, penalty, pool storage. 1. introduction sources of uncertainties are contained in current approach for criticality calculations . most of these uncertainties are considered in uncertainty evaluation, even so some of them remain and are not evaluated. it could possibly lead to non–conservative evaluation and even to violation of safety limits. computational code for criticality calculations is validated based on specific types of critical experiments. thereafter is said that code gives correct believable results. this paper focuses to selection of experiments used for validation, because main omitted source of uncertainties is contained there. one purpose for validating codes at department of reactor physics (nri rez) are calculations of spent fuel storages. code is usually validated using critical experiments with fresh fuel. however when code is applicated to spent fuel storage there is not considered effect of disparity between used experiments and newly suggested spent fuel storage. this brings potential risk and non–conservativeness described above. goal to be achieved was to evaluate effect of disparity and prepare instructions for validating code application on spent fuel systems. whole research was done by means of scale [1] with advantage of it’s module tsunami (uses keno monte carlo code, will be described later). 2. current process of validation and uncertainty evaluation process of validation and uncertainty evaluation in criticality calculations is currently carried out this way: (1.) firstly set of critical experiments for validation is chosen – selection usually depends on which experiments are available • for each experiment (i) are known both keff – measured (ke,im ) and calculated (ke,ic ) (2.) then biases bi for each experiment (i) are calculated (a) bi = ke,im −ke,ic , if ke,ic < ke,im (b) bi = 0, if ke,ic >= ke,im (3.) average bias is calculated from bi (4.) model of new system, which is newly designed (called application), is calculated → kac ±σkac • where σkac is statistic error of keno-vi monte carlo calculation (5.) finally sum of kac + 2σkac + bias + 2σbias (1) must be smaller than required safety limit (usually set by legislation) as was already pointed out above, in this procedure disparity between experiments and applications is not considered. it affect mainly two parts of the procedure: (1.) experiments for evaluation of bias are chosen without consideration of similarity between the application and the experiment (2.) in the final equation there is no term expressing and quantifying potential deviation caused by disparity between experiment and application 30 http://dx.doi.org/10.14311/app.2018.19.0030 http://ojs.cvut.cz/ojs/index.php/app vol. 19/2018 uncertainty evaluation in criticality calculations both of these aspects can be resolved using tsunami. 3. tsunami tsunami is part of complex computational code scale and its name stands for tools for sensitivity and uncertainty analysis methodology implementation. complete documentation can be found in [1]. when calculating systems parameters using tsunami-3d-k6 (monte carlo code keno is used for these calculations), one of the possible outputs is the sensitivity data file (.sdf). these contains sensitivity informations (how much keff changes because of change of cross sections – δk δσ ) for each nuclide, energy group and reaction. from these files tsunami-ip (indices and parameters) calculates amount of similarity between chosen systems. several similarity coefficient exist, which are calculated different ways based on sensitivities and uncertainties of cross section determination. most important coefficient, based on both sensitivity and uncertainty is ck. it expresses overall similarity of two systems (detailed description can be found in [1]). coefficient ck is normalized such way, that similar systems have ck = 1 and totally dissimilar systems have ck = 0. currently similarity with ck > 0.95 is considered sufficient. second important value is penalty. it is calculated based on both sensitivity and uncertainty. penalty expresses in % ∆kk uncertainty of calculating application system based on dissimilarity between experiments and applications. when calculating penalty it is possible to choose which experiments should or should not be used for penalty assessment. parameter cvalue can be set to ensure that only experiments with ck > cvalue will be used for penalty evaluation. using cvalue is necessary – it is not required to deduct informations from totally dissimilar systems. each system has maximum value for penalty, which responds with relative standard deviation of keff due to cross-section covariance data. usually in results of various systems calculations observed maximal penalty was around 0.6% ∆kk , which corresponds with [2]. value of maximum penalty is included as “relative standard deviation of keff due to cross-section covariance data” in output, when system is calculated with tsunami-3d-k6 module from scale. diagram of entire process of calculations and analysis using tsunami is shown in figure 1. 4. comparing system similarity and validation quality for various types of experiments for further usage determination of behaviour of similarity coefficient was needed. firstly several pincell systems with varying parameters were tsunami-3d-k6 tsunami-ip penalty c k .sdf files max. penalty cvalue applications inputs experiments inputs figure 1. diagram of evaluation process using tsunami. compared for this purpose. later, spent fuel wet storages as applications and different types of experiments were examined. 4.1. simplified systems infinite systems of fuel rods were modelled. for comparison some parameters within systems were changed: (1.) grid type (2.) enrichment (3.) pitch (4.) amount of absorber (h3bo3) (5.) isotopic composition of fuel due to burnup. in each case only one parameter was changed, others remained nominal, which means: • enrichment 4% • pitch between fuel rods 0.61 cm • 4 g/kg of h3bo3 in moderator. when comparing triangular and square grid the moderator-to-fuel ratio for compared systems remained unchanged. multiples from 0.8 to 2.5 of nominal pitch between fuel rods were used in infinite grids. for each multiple of nominal pitch ck between triangular and square grid was bigger than 0.99, thus the experiments with square grid can be used for validation for calculating triangular grids and same the other way. after comparison of square and triangular pitch and proving their similarity, all further comparisons were done for triangular grid. comparison of similarity between systems with different enrichments is displayed in figure 2. system with enrichment 4% was used as the experiment and 31 matěj šikl, františek havlůj acta polytechnica ctu proceedings 0 5 10 15 20 0.2 0.4 0.6 0.8 1 enrichment [%], original value 4.0 c k 0 0.2 0.4 0.6 0.8 1 p en a lt y [% d k / k ] c k c k=0.95 penalty figure 2. evaluation of similarity between systems with different enrichments. 0.4 0.6 0.8 1 1.2 0.7 0.8 0.9 1 pitch [cm], original value 0.61 c k 0 0.2 0.4 0.6 0.8 1 p en a lt y [% d k / k ] c k c k=0.95 penalty figure 3. evaluation of similarity between systems with different pitches. applications had from 0 to 20% enrichment. in figure 2 it can be seen that variation in enrichment from 1.5 to 8% has still ck > 0.95. that seems to be a sufficient range for a typical application. figure 3 shows comparison between systems with different pitches. grid with nominal pitch 0.61 cm was taken as the experiment. as can be seen in figure 3, systems with changes in pitch smaller than 0.1 cm are still sufficiently similar. value of 0.1 cm seems to be small variability, but in fact is enough, because experiments will be typically available for identical or similar types of fuel assemblies as in the applications. boron acid amount in moderator was next parameter changed. system with 4 g/kg boron acid in moderator was taken as the experiment. figure 4 shows that comparison and it can be seen that ck remains bigger than 0.99 for all applications. this fact is important, because it means it is possible to use 0 2 4 6 8 10 0.95 0.96 0.97 0.98 0.99 1 1.01 1.02 amount of h3bo3 [g/kg], original value 4.0 c k 0 0.2 0.4 0.6 0.8 1 p en a lt y [% d k / k ] c k penalty figure 4. evaluation of similarity between systems with different amount of h3bo3 in moderator. 0 10 20 30 40 50 60 0.5 0.6 0.7 0.8 0.9 1 burn-up mwd/kgu c k 0 0.1 0.2 0.3 0.4 0.5 p en a lt y [% d k / k ] c k penalty figure 5. comparison of irradiated fuel with fresh uo2 fuel. experiment without boron for application including boron with no big impact. later, systems with fuel rods in different degrees of irradiation were calculated and compared. to get closer to reality systems with fresh fuel only were used as experiments, because no experiments with irradiated fuel are included in evaluated databases of critical experiments. applications are systems with isotopic composition obtained with origen, input for origen was fresh uo2 fuel. firstly irradiated fuel applications were compared with fresh uo2 fuel as the experiment and is shown in figure 5. in figure 5 can be seen that if ck will be required bigger than 0.95, only spent fuel with irradiation smaller than approximately 10 mwd/kgu are sufficiently similar. 32 vol. 19/2018 uncertainty evaluation in criticality calculations element atom density pu-238 7.39e-06 pu-239 4.10e-03 pu-240 5.61e-04 pu-241 4.38e-05 pu-242 1.70e-05 u-235 1.20e-04 u-238 1.69e-02 o 4.37e-02 am-241 3.43e-05 table 1. fresh mox material composition. 0 10 20 30 40 50 60 0.4 0.5 0.6 0.7 burn-up mwd/kgu c k 0.34 0.36 0.38 0.4 0.42 0.44 0.46 0.48 0.5 p en a lt y [% d k / k ] c k penalty figure 6. comparison of irradiated fuel with fresh mox fuel. secondly fresh mox fuel was used as the experiment. mox material composition used in calculations follows in table 1. mox was chosen because during burnup actinides are created (most important is pu) and in mox fuel pu and am isotopes are contained. output of comparison is shown in figure 6 and interesting progress in ck can be observed. applications with small burnup are not similar with experiment (ck < 0.5), but later with burnup raises to ck = 0.7. this progress led to comparing applications (irradiated fuel) with both experiments (uo2 and mox) in one time. this comparison is shown in figure 7. figure 7 brings one essential fact: ck > 0.95 is currently unfulfillable requirement. 4.2. realistic systems later more realistic systems were calculated, after first simplified comparison was completed. experiments were taken from international handbook of evaluated criticality safety benchmark experiments (icsbep, [3]) database. this database collects nearly 5000 critical experiment and benchmarks from nuclear facilities around the world. 0 10 20 30 40 50 60 0.75 0.8 0.85 0.9 0.95 1 burn-up mwd/kgu c k 0 0.1 0.2 0.3 0.4 p en a lt y [% d k / k ] c k penalty figure 7. comparison of irradiated fuel with fresh uo2 and mox fuel experiments together. systems with low, intermediateand highlyenriched uranium, mox and other types of fuel are included. for some experiments are available input files (scale or mcnp) and for some are available .sdf files. for following comparison .sdf files from there were used, even though they were produced with another version of nuclear data libraries. this was done to save computational time and for possibility for using hundreds of experiments. applications were systems representing an infinite lattice of dukovany nuclear power plant wet spent fuel storage cells with varying fuel burnup. in this wet spent fuel storage two racks are located – the upper and the lower, where the lower one has boron metal sheets placed between fuel assembly positions. for the application a fuel assembly gd2m with mean enrichment 4.38% with burnup from 0 to 60 mwd/kgu was chosen. maximum possible penalty for the applications in question ranges from 0.4 to 0.6% ∆kk . as it was said, variable types of experiments were available and comparison was performed with various combinations of experiments. results of comparison were poor for application with small burnup when experiments were mox and for applications with high burnup when experiments were fresh low enriched uranium fuel, which corresponds with results from section 4.1. therefore result for comparison with mix of available experiments with low enriched uranium and mox will be discussed. first, comparison with variable cvalue (different for each application) was performed. in the first run ck were calculated and from these the cvalue for each application was set as 0.75 multiple of the highest ck achieved for individual application. value 0.75 was chosen because of two reasons. firstly experiment filtration is necessary, so only adequate system remains and are used in analysis. secondly acceptable amount of experiments is needed, so not to filter majority 33 matěj šikl, františek havlůj acta polytechnica ctu proceedings grid, mwdkgu cvalue penalty [% ∆k k ] cvalue passed bottom, 0 0.68 3.24e-2 402 / 726 bottom, 15 0.62 3.88e-2 489 / 726 bottom, 30 0.58 6.48e-2 489 / 726 bottom, 45 0.59 1.52e-1 219 / 726 bottom, 60 0.57 2.09e-1 228 / 726 upper, 0 0.70 2.24e-2 359 / 726 upper, 15 0.64 5.13e-2 395 / 726 upper, 30 0.59 1.05e-1 291 / 726 upper, 45 0.58 2.02e-1 164 / 726 upper, 60 0.55 2.79e-1 166 / 726 table 2. comparison of wet spent fuel storage applications with low enriched uranium and mox fuel experiments, variable cvalue. grid, mwdkgu c max k penalty [% ∆k k ] cvalue passed bottom, 0 0.90 3.28e-2 401 / 726 bottom, 15 0.82 1.70e-1 367 / 726 bottom, 30 0.77 1.56e-1 139 / 726 bottom, 45 0.79 1.57e-1 139 / 726 bottom, 60 0.76 2.17e-1 82 / 726 upper, 0 0.93 2.24e-2 358 / 726 upper, 15 0.85 5.37e-2 273 / 726 upper, 30 0.79 1.11e-1 67 / 726 upper, 45 0.78 2.18e-1 64 / 726 upper, 60 0.73 3.09e-1 7 / 726 table 3. comparison of wet spent fuel storage applications with low enriched uranium and mox fuel experiments, cvalue=0.7 fixed. of all experiments is required. multiplication by 0.75 seemed to fulfil both requirements. the results of comparison of both grids and various levels of irradiation as applications and both together low enriched uranium and mox fuels as experiments are recorded in table 2. in comparison in table 2 for each application penalty value is smaller than maximum penalty. standard deviations of the penalties were approximately 2 orders smaller than the penalty values. as expected (because of choice of multiplicative factor) for each application some experiments passed cvalue test. on the other side cvalue for some applications was approaching 0.5, which is unacceptably low value, hence another system for setting cvalue had to be used. that system has to be independent on application’s maximal ck. paper [4] deals with requirements cvalue should fulfil. in accordance with that paper cvalue=0.7 was fixed in further investigations. results can be seen in table 3. in table 3 standard deviations of the penalties were approximately 2 orders smaller than the penalty values for each application. last comparison (results in table 3) still shows good results for penalty, its value is multiply smaller than value of maximum penalty. number of experiments passed through cvlaue test is on one side small, on the other side not all of the experiments were filtered. this process of calculating and validating seems to be best of all tried. for spent fuel applications mox and low enriched uranium experiments will be needed. in next section optimized recommended approach for validating will be described. 5. recommended optimized process based on research realized and knowledge from papers [4] and [5] an optimized validation and uncertainty evaluation in criticality calculations process will be recommended. the required value 0.7 for cvalue test is based on [4], requirement for at least 40 critical experiments passing cvalue test is based on [5]. as it was said in introduction, main issues in the current approach for validation are the way for choosing experiments for determining bias and the neglection if dissimilarity between experiments used in validation and applications. both of these can be resolved using tsunami. firstly, tsunami-ip analysis should be done for experiments considered for use in bias evaluation. the choice of experiments should be: • if at least 40 experiments with ck > 0.7 to application will be found → these experiments should be used for bias determination • if less than 40 experiments with such similarity will be found → 40 experiments most similar (with highest ck) should be used for determinating bias. for determining the value of penalty the following approach should be used: • if at least 40 experiments with ck > 0.7 to application will be found → value of penalty obtained from tsunami can be used • if less than 40 experiments with such similarity will be found → value of maximum penalty should be used. with this approach, the equation 1 is changed to kac +2σkac +bias +2σbias+penalty+2σpenalty < usl, (2) where usl stands for upper safety limit. 6. conclusion the issues in the current approach for validating and uncertainty evaluation in criticality calculation were pointed out and goal to resolve these was set. when evaluating criticality calculations, many sources of uncertainties are included, but influence of dissimilarity between system is currently neglected. this influence could be evaluated using tsunami methodology. for these purpose tsunami was described and calculations were performed. from these calculations possible ranges of individual parameters for sufficient similarity were determined. 34 vol. 19/2018 uncertainty evaluation in criticality calculations moreover recommendation for selection of critical experiment used for validation was found. all of these experiences were used to produce a recommended approach described in section 5, where easy to follow guide is placed. during the whole research and preparing recommended approach an effort to optimization was carried out, hence recommended approach has not too strict requirements while requesting all important inputs and being conservatively safe. correctly performed procedure can minimize deviation due to disparity of systems to level of calculation statistic error. the goal of preparing a complex approach for validating and uncertainty evaluation in criticality calculation was fulfilled and it was done in such way that it is possible to prepare automatic script for processing. references [1] b. t. rearden, m. a. jessee. scale code system, ornl/tm-2005/39, version 6.2.2. oak ridge national laboratory, oak ridge, tennessee, 2017. [2] m. klein, l. gallner, b. krzykacz-hausmann, et al. influence of nuclear data uncertainties on reactor core calculations. kerntechnik 76:174–178, 2011. [3] international handbook of evaluated criticality safety benchmark experiments, 2016. nea/nsc/doc(95)03, oecd, nea no. 7328. [4] b. rearden, m. williams, m. anderson jessee, et al. sensitivity and uncertainty analysis capabilities and data in scale. nuclear technology 174:236–288, 2011. [5] b. l. broadhead, b. rearden, c. hopper, et al. sensitivityand uncertainty-based criticality safety validation techniques. nuclear science and engineering 146:340–366, 2004. 35 acta polytechnica ctu proceedings 19:30–35, 2018 1 introduction 2 current process of validation and uncertainty evaluation 3 tsunami 4 comparing system similarity and validation quality for various types of experiments 4.1 simplified systems 4.2 realistic systems 5 recommended optimized process 6 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0052 acta polytechnica ctu proceedings 22:52–56, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app experimental analysis focused on the material characteristics of the cement-based polymer-modified mortars barbara kucharczyková∗, hana šimonová, romana halamová, dalibor kocáb, martin alexa brno university of technology, faculty of civil engineering, veveří 331/95, 602 00 brno, czech republic ∗ corresponding author: barbara.kucharczykova@vutbr.cz abstract. the article deals with the experimental analysis focused on the development of the physical and mechanical characteristics of the cement-based polymer-modified mortars (pcm) during ageing. two different commercial products commonly used for similar in-situ applications were used for the experiment. the results of the shrinkage, elastic, fracture and strength parameters determined within the time interval from 3 days to 2 years of ageing are summarized and discussed in the article. the performed experimental analysis showed different behaviour of tested pcms. the most significant differences were observed at the age of 90 days when one of the tested pcm showed a substantial decrease in most of investigated characteristics. keywords: polymer-modified mortar, cement, shrinkage, fracture, modulus of elasticity, strength. 1. introduction cementitious mortars modified by polymers, commonly known as polymer-cement mortars (pcm), find wide use in civil engineering especially in the rehabilitation and protection of both concrete and masonry structures [1, 2]. pcm have a monolithic matrix that uniformly combines the matrix of an organic polymer with the matrix of a cement gel [3]. thanks to the homogenization of the cement matrix with the polymer, these mortars receive an improvement in some properties – e.g. workability, tensile strength, adhesion, or chemical corrosion resistance [4, 5]. cementitious mortars can be modified by various polymers; e.g. latex, redispersible polymer powders, liquid resins, monomers, or acrylic and epoxy emulsions. polymers and polymer-modified materials can nowadays be considered an important component of modern, sustainable civil engineering [6]. applications where pcm are best used as well as how they are to be technologically implemented are regulated by documents and technical manuals supplied with each product; however, these instructions can often be rather vague. if pcm are wrongly applied, they may fail to perform and put the functionality of the entire rehabilitation system in danger [7]. differences in the behaviour of pcm can also be expected with varying quality and type of substrate. they may behave differently in combination with e.g. high-quality concrete compared to non-fired bricks. it is also important to properly treat the surface of the substrate (removing dust, wetting and saturating with water) as well as curing the final pcm surface once exposed to the environment. in order to understand the behaviour of pcm and their use it is necessary to know the value of compressive and tensile strength at the age of 28 days as well as the progress of their other material properties, such as volume changes or crack propagation resistance over a long period of time. 2. experimental part the main aim of the performed measurements was to determine the development of the selected material characteristics of the pcms during the ageing and show the specific behaviour which was observed during testing. to simulate inadequate curing conditions which can occur during in-situ applications, all tested specimens were not treated at all during the whole time of ageing and were stored in laboratory conditions with temperature of (21 ± 2) °c and relative humidity of (60 ± 10)%. for the first 72 hours they were stored in the moulds with uncovered upper surface. after demoulding all surfaces were left to dry freely. the results presented in this paper were partially presented in [7] and they are extended herein by the results of the tests performed at the specimens’ age of approx. 2 years. 2.1. materials two fine-grained polymer-modified mortars based on the portland cement were prepared for the experiments. because the commercial products were used, the trademarks and the composition of these products cannot be specified in the paper. only general information taken from the technical sheets were used for the specification of the materials. note, that both mortars are commonly used for similar in-situ applications. for the purpose of the results evaluation, the pcms were designated as “v” and “vii” sets. the specimens of the “v” set were made of a twocomponent mixture where a liquid component was an 52 http://dx.doi.org/10.14311/app.2019.22.0052 http://ojs.cvut.cz/ojs/index.php/app vol. 22/2019 experimental analysis focused on the material characteristics. . . aqueous copolymer dispersion and a powder component contained a mixture of portland cements and mineral fillers. the specimens of the “vii” set were made of a single-component powder mix, containing among others silica sand, portland cement, microfibers including plasticisers and polymers, which was mixed with the prescribed dosage of water to prepare a fresh mortar [7]. both pcms were prepared using a handheld mixer. the mixing process, including the mixing time, was proceeded in compliance with the technical sheet of each product. two types of the test specimens were manufactured for the purpose of experiment. the specimens with dimensions of 60 × 100 × 1000 mm were used for the measurement of shrinkage process. remaining characteristics were determined using the prismatic specimens with dimension of 40 × 40 × 160 mm. all experimental results were evaluated for the specimens’ ages of 3, 28, 90 and 730 days and were expressed by the mean value and sample standard deviation calculated from three and six independent measurements in the case of shrinkage and fracture tests respectively. 2.2. testing techniques 2.2.1. shrinkage the shrinkage process was determined using the shrinkage drains with a movable head [8] which enabled to start the measurement of the relative length changes very early after the fresh-state material was placed into the measuring moulds. the inner surfaces of the shrinkage drains with the internal dimensions of 60 × 100 mm in cross-section and 1000 mm in length were coated with a 2 mm thick polyethylene foam mat to avoid a friction between the mould and test specimen during the measurement. in this way the relative length changes were measured along the central axis of the specimen using an inductive sensor leaning against the movable head of the mould during about 72 hours. special markers were also embedded into the upper surface of the test specimens during their manufacturing to facilitate long-term measurement of relative length changes after the specimens were removed from the drains. the arrangement of the measurement is shown in fig. 1. refer to [9] for the test procedure details. 2.2.2. dynamic modulus of elasticity each specimen was tested for the natural frequency of longitudinal vibration, see fig. 2. the vibration was produced by a mechanical impulse generated by an impact hammer and the natural frequency was measured using an oscilloscope handyscope hs4 with an acoustic emission sensor. the dynamic modulus of elasticity was then calculated from this measured natural frequency according to the standard [10] using the equation: edyn = 4 · l2 · f 2l · d, (1) where edyn is the compressive dynamic modulus of elasticity in [mpa], l is the specimen length in [m], fl is the natural frequency of longitudinal vibration measured in [khz] and d is the material’s bulk density in [kg/m3]. 2.2.3. fracture test the fracture characteristics were determined based on the results of the three-point bending test of prismatic specimens with dimensions of 40 × 40 × 160 mm provided with an initial notch located in the middle of the specimens’ length. the depth of the notch corresponded in all cases approximately to the 1/3 of the specimen’s height. the span length was set to 120 mm. a mechanical testing machine fp10/1 was used for testing which enabled the displacement increment loading of the test specimens with a rate of 0.02 mm/min. the loading force (f ) and deflection (d) measured in the middle of the span length were continuously recorded into the data logger during the test. based on the f − d diagrams the elastic modulus was calculated from the initial part of the diagram. the effective fracture toughness was determined using the effective crack model and the work of fracture and the consequent specific fracture energy was calculated using work-of-fracture method [11, 12]. because of the loss of stability, which occurred during the loading, the value of the work of fracture w ∗f was determined as an area under the proper part of f − d diagrams before stability loss occurred. refer to [13] for details about the determination of above mentioned characteristics. the informative value of compressive strength was determined on the specimens’ fragments after the fracture tests were finished. arrangement of the fracture and compressive test is shown in fig. 2. 3. results and discussion the results of performed measurements represented by the average values (markers on the lines) and sample standard deviations (introduced in the table) for the specific age of specimens are shown in fig. 3 – fig. 5. the results displayed in fig. 3 show development of the compressive strength (fc) and dynamic modulus of elasticity (edyn). different trend and also different absolute values of these characteristics were recorded for the set “v” and “vii”. in general, the “vii” set shows higher values of fc up to the age of 90 days in comparison with the set “v”. on the other hand, a visible decrease in this characteristic was recorded at the age of 730 days, when the magnitude of the fc is lower than the one at the age of 28 days and is very similar to the fc value determined for “v” set at the same age. the curve of edyn values determined for “v” set shows gradual increase during monitored time-interval, whereas the “vii” set values started to decrease at the age of 28 days. in general, it can be stated that increase of edyn values was insignificant during the monitored time-interval. the results of 53 b. kucharczyková, h. šimonová, r. halamová et al. acta polytechnica ctu proceedings figure 1. arrangement of the relative length changes measurement. figure 2. arrangement of the fracture test (left), compressive strength test (middle) and resonance test (right). figure 3. development of the compressive strength (left) and the dynamic modulus of elasticity. 54 vol. 22/2019 experimental analysis focused on the material characteristics. . . figure 4. development of the modulus of elasticity determined from f − d diagrams (left) and the shrinkage process. figure 5. development of the effective fracture toughness(left) and the specific fracture energy. both characteristics are accompanied by relatively high variability. in the case of the “vii” set, interesting trends are observed in the development of the characteristics obtained from the fracture tests (fig. 4 – left and fig. 5). a steep decrease in the values of elastic modulus (ef r), effective fracture toughness (kice) and specific fracture energy (gf ) occur after the age of 28 days. the values of ef r, kice and gf decrease by about 70% within the age of 28 and 90 days. one of the explanation of this steep decrease can be found in the shrinkage curve (fig. 4 – right) which shows a steep increase in the shrinkage values which reached almost the maximum within the 28 day of specimens’ ageing. the results are also accompanied by a high variability. the tests performed at the age of 730 days show an increase in the values of all fracture characteristics within the 90 and 730 days. the values increased approx. 2.5 times, 3 times and more than 4 times in the case of ef r, kice and gf respectively. this increase is supposedly associated with a self-healing ability of the material. results of fracture characteristics for “v” set do not show any dramatic decline of the fracture characteristics. the ef r values slowly increase during whole time, values of remaining characteristics (kice, gf ) increase up to the age of 28 days. after that both values decrease slightly. the gf value shows a gradual decrease up to the age of 730 days. the shrinkage curve (fig. 4 – right) of “v” set differs slightly from the curve of “vii” set. the shrinkage curve determined for “v” set is not so steep within the 3 and 90 days of ageing and the final shrinkage value is about 25% lower. the results show also much lower variability. 4. conclusions despite the fact that both tested pcms are used for similar in-situ applications, the performed experimental analysis, focused on the long-term development of the selected material characteristics, showed different behaviour of particular pcms. the greatest differences were observed at the age of 90 days when, in the case of “vii” set, most of investigated characteristics significantly decreased in their values. on the other hand, the tests performed at the age of 730 days for “vii” set showed some self-healing ability of this material which led to the re-increasing of the fracture characteristics especially. the shrinkage curves, determined on the test specimens intentionally exposed to the free desiccation during whole monitored time-interval, showed a steep increase in the shrinkage values for both pcms. this fact accentuates the necessity to use appropriate curing conditions during theirs in-situ applications. important is also the fact that most results for both pcms were accompanied with a high variability. 55 b. kucharczyková, h. šimonová, r. halamová et al. acta polytechnica ctu proceedings acknowledgements this paper has been written as part of the project no. ga17-14302s “experimental analysis of the early-age volume changes in cement-based composites”, supported by the gao czech science foundation and project no. fast-j-18-5516 “determination of modulus of elasticity in the early phase of maturing of cement composites and their thermal expansion”. references [1] w. malorny, m. plath. investigations on properties determining durability of novel pcc. in proceedings of the conference on the rehabilitation and reconstruction of buildings crrb 2012, vol. 688 of advanced materials research, pp. 130–138. trans tech publications, 2013. doi:10.4028/www.scientific.net/amr.688.130. [2] o. karel, j. kubanek, p. schmid, i. rozsypalová. sanace panelových bytových domů pomocí dodatečně vkládané helikální výztuže – vývoj, alternativy, zkoušení. beton tks 03:24–27, 2017. [3] y. ohama. handbook of polymer-modified concrete and mortars properties and process technology. park ridge, n.j: noyes publications, 1995. [4] l. aggarwal, p. thapliyal, s. karade. properties of polymer-modified mortars using epoxy and acrylic emulsions. construction and building materials 21(2):379–383, 2007. doi:10.1016/j.conbuildmat.2005.08.007. [5] p. łukowski. polymer-cement composites containing waste perlite powder. materials 9:839, 2016. doi:10.3390/ma9100839. [6] l. courard, j.-f. lenaers, f. michel, a. garbacz. saturation level of the superficial zone of concrete and adhesion of repair systems. construction and building materials 25:2488–2494, 2011. doi:10.1016/j.conbuildmat.2010.11.076. [7] b. kucharczyková, h. šimonová, z. keršner, d. kocáb. shrinkage, weight loss and fracture parameters of selected polymer-modified cement mortars during ageing. transactions of the všb: technical university of ostrava, civil engineering series 17(01):122–130, 2017. doi:10.1515/tvsb-2017-0014. [8] schleibinger gerate. schleibinger testing systems: ultrasonic setting measurement. http://www.schleibinger.com, 2018. [9] čsn 73 1372 non-destructive testing of concrete – testing of concrete by resonance method. standard, úřad pro technickou normalizaci, metrologii a státní zkušebnictví, 2012. [10] b. kucharczyková, l. topolář, p. daněk, et al. comprehensive testing techniques for the measurement of shrinkage and structural changes of fine-grained cement-based composites during ageing. advances in materials science and engineering 3:1–10, 2017. doi:10.1155/2017/3832072. [11] rilem tc-50 fmc recommendation, determination of the fracture energy of mortar and concrete by means of three-point bend tests on notched beams. materials and structures 18(4):287–290, 1985. doi:10.1007/bf02472918. [12] v. veselý. the role of process zone in quasi-brittle fracture. habilitation thesis, university of technology, faculty of civil engineering, institute of structural mechanics, brno, 2015. [13] p. rovnaník, h. šimonová, l. topolář, et al. carbon nanotube reinforced alkali-activated slag mortars. construction and building materials 119:223–229, 2016. doi:10.1016/j.conbuildmat.2016.05.051. 56 http://dx.doi.org/10.4028/www.scientific.net/amr.688.130 http://dx.doi.org/10.1016/j.conbuildmat.2005.08.007 http://dx.doi.org/10.3390/ma9100839 http://dx.doi.org/10.1016/j.conbuildmat.2010.11.076 http://dx.doi.org/10.1515/tvsb-2017-0014 http://www.schleibinger.com http://dx.doi.org/10.1155/2017/3832072 http://dx.doi.org/10.1007/bf02472918 http://dx.doi.org/10.1016/j.conbuildmat.2016.05.051 acta polytechnica ctu proceedings 22:52–56, 2019 1 introduction 2 experimental part 2.1 materials 2.2 testing techniques 2.2.1 shrinkage 2.2.2 dynamic modulus of elasticity 2.2.3 fracture test 3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0055 acta polytechnica ctu proceedings 13:55–60, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app micro-mechanical performance of concrete used as recycled raw material in cementitious composite vladimír hrbeka, b, ∗, veronika koudelkováa, b, zdeněk prošeka, c, pavel tesáreka a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b institute of theoretical and applied mechanics as cr, v.v.i., prosecká 76, 190 00 prague 9, czech republic c university centre for energy efficient buildings of czech technical university in prague, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: vladimir.hrbek@fsv.cvut.cz abstract. the reduction of industrial pollution is recently one of main goals over all fields. in civil engineering, re-cycling of structural waste provides wide opportunity contributing this effort. this paper focus on re-use of concrete waste, which after further processing can be used in new constructions as partial supplement to the mixture. to investigate the impact of re-cycled concrete addition, it is necessary to determine mechanical and structural parameters of individual phases in the “raw” material. for this purpose, grid indentation and scanning electron microscopy with energy-dispersive x-ray spectroscopy (sem, edx) are combined to determine properties of concrete sample. keywords: re-cycled concrete, grid indentation, sem, edx, image analysis. 1. introduction the waste material accumulation became worldwide environmental and economical problem. development of new effective and thrifty methods for re-use of waste products lowers the number of waste deposits while saving natural sources and environment. for example, in 2012, the countries of european union produced over 2.5 billion tones of waste and managed to recycle barely 10% of this amount. in case of czech republic, over 23 million tones of waste were produced and slightly over one-fifth recycled (statistics according to environmental data center on waste). one of main problems with incorporation of waste materials in civil engineering (especially in constructions) is mistrust of designers and developers toward recycled construction waste. despite lower price of these materials (compare to traditional raw materials), final cost can be paradoxically higher due to the transportation [1–3]. the solid phase of concrete is made of coarse and fine aggregate (65 to 70%) and cement paste. typically, when recycled, coarse aggregate (30 to 50% of raw material) is to be reused, which leaves majority of the concrete as secondary waste product [4]. complete recycling of concrete (concretely use of highspeed milling) can turn this waste into fine powder (fraction < 1 mm) with potential usage as micro-filler (fine ground hydration products and fine aggregate) and partial cement substitute (non-hydrated clinker in cement paste). nevertheless, the processing itself is complicated and bears addition financial costs. beyond that, due to the lack of studies and short experience of civil engineering practice with such fine powdered material, the implementation of recycled secondary waste material is not possible on large scale. m. chen et al. [5] used the recycled concrete fine powder in their study as full substitute of pulverized limestone in asphalt mixture. the results of x-ray diffraction and scanning electron microscopy (sem) showed higher content of sio2, lower caco3 content and rougher partical surface in recycled concrete fine powder, compare to of pulverized limestone. when used as filler in asphalt mixture, the water sensitivity of samples reduced while mechanical properties in high temperatures and fatigue persistence increased. however, mechanical properties in low temperature slightly decreased. in case of cement based mixtures, recycled concrete fine powder serves much as a micro-filler, which acts as coarse aggregate on the mezoand micro-scale of the material. unfortunately, studies about recycled micro-fillers are not yet carried out, although the idea seems promising as replacement of current microfilling agents [6]. moreover, the potential of nonhydrated clinker present in the powder was so far vastly ignored. y. j. kim et al. [7] presented in his study features of cement paste modified by recycled concrete fine powder as a partial replacement for a binder. according to the results, workability in mixture’s fresh state decreased with increasing level of cement compensation and the hydration process was delayed up to 2 hours. in case where 45% of cement was replaced by recycled concrete fine powder, the mechanical com55 http://dx.doi.org/10.14311/app.2017.13.0055 http://ojs.cvut.cz/ojs/index.php/app v. hrbek, v. koudelková, z. prošek, p. tesárek acta polytechnica ctu proceedings pressive strength reduced to 30% while absorptivity gain reached almost 70%. the conclusion of this study proposed limitation of the replacement to 15%. m. lidmila a k. šeps [8–10] used in their studies pulverized concrete recycled from railway sleepers. the product of recycling showed low reactive potential very likely due to very fine milling of the concrete, i.e. unfolding of non-hydrated clinkers. as the level of non-hydrated clinker is low, the material is to be possibly used for strengthening of subsoil under railway structures [11]. other potential utilization, in small-scale, can be found in cementitous composite material. the knowledge of the material on the micro-scale is thus crucial to understand the composite’s behavior on higher scale levels. with such level of investigation, it is possible to develop up-scaled prediction of macro-mechanical properties, as well as optimize the efficiency of used fine powdered concrete recycle product. this study is focused on investigation of micro-mechanical and micro-structural performance of the in-put “raw” material, i.e. concrete sleepers intended for recycle. the output would serve as comparative analysis to further planed research of the recycled pulverized concrete. 2. material on of raw materials used for the recycled pulverized concrete are railway sleepers type pb2 and sb8 (used in this study, fig. 1a). it the first step of the recycling process, crude recycled concrete gravel with dimension 0–32 mm (fig. 1b) arises from the kibble of the sleepers and the metallic parts and reinforcement is separated. after selection of finer concrete gravel (fraction 0–16 mm, fig. 1c), the material is milled in two steps to the final product (recycled pulverized concrete, fraction 0–1 mm). the fine milling of the material was delivered by company lavaris ltd. (procedure protected by patents act) using the high speed milling technique. for a purpose of the investigation, piece of crude concrete gravel (fig. 1b) was embedded in epoxy resin and let to harden for approx. 24 hours. a part of the specimen was than severed (approx. thickness 15 mm) and silica-carbon papers and diamond suspensions were used for grinding and polishing of the specimen surface. this technique ensured adequate surface roughness of the sample for both indentation and sem investigation. the coordination mark was inscribe on the polished surface, serving as origin for further orientation. this enabled positioning of both investigation techniques in the very like area of the sample. 3. investigation procedure the investigation of the crude recycled concrete consisted of three main techniques, instrumental indentation, scanning electron microscopy combined with (a) . railway sleepers, type pb8. (b) . crude recycled concrete (0 32 mm) (c) . “raw” recycled concrete (0 16 mm) figure 1. products of concrete recycling procedure. electron probe microanalysis (sem epma) and back scattered electron micrographs (sem bse) image processing. as a result, combination of micromechanical properties, phases chemical composition and material structure enabled proper description of the material. 3.1. scanning electron microscopy the scanning electron microscopy (sem) investigation was performed in mira ii lmu (tescan corp., brno) on polished specimens coated with thin layer of carbon necessary to ensure proper conductivity of the surface. working distance of the microscope was set closely to 15 mm and accelerating voltage was set to 15 kv – to provide good signal. chemical composition (in weight percentage) of each phase characterized by the same level of grey scale was determined using epma based on the detection of the x-rays by energy dispersive x-ray detector (edx, bruker corp., berlin) on several point of each position. due to the time-demands of technique such as phase determination by epma, image analysis of sem bse micrographs was used to obtain phase volumetric representations in viewed field, i.e. material structure. the micrographs were acquired at different magnifications. the approximate area of indentation was investigated at first at the magnification 600× to gain adequate overview. than the area was scanned with higher magnification (above 1200×) for detailed study of indented positions in several images. the sem micrographs acquired with back scattered electron detector (bse) provide information about distribution (and chemical composition) of different phases due to different spectrum of the back scattering coefficient. 56 vol. 13/2017 re-cycled concrete 3.2. quasi-static nano-indentation the technique of nano-indentation (ti 7500 series, hysitron inc.) was selected for evaluation of microscale elastic properties of the material. its principal is based on dependency of probe propagation with respect to the recorded force. the force-displacement record is further further processed and the mechanical properties are calculated from the unloading part of the record. incorporated errors of the measurement, such as creep and visco-elasticity of measured phases, is avoided by appropriate test setup [12–18]. for purpose of this study, the load function kept for further described indentation similar, i.e. the “loading” and “unloading” segments lasted 5 seconds each with an in-between 60 seconds “holding” time segment. the measurement was performed in two major steps (level i and ii) to achieve efficient investigation of phases on different material levels. as an early step, a load-controlled quasi-static grid indentation was performed with an applied force of 10 mn (almost maximum capacity of the transducer) to obtain route-identification of the material. the indents were each separated by 25 µm in 10 by 10 raster, which allowed both micro-mechanical evaluation of phases in large range and estimation of further used indentation depth. however, such setting does not meet the grid-indentation specifications of heterogeneous composite [19–22], particularly does not guarantee consistent half-space influenced by each indent. the first step (level i) was thus selected as displacement-driven quasi-static grid measurement with each indent driven into 150 nm depth in two positions. the grid was constructed of 21 by 21 indents pattern with 5 µm spacing. the setting ensured equal indented half-space volume of material represented on the sample surface by circle with 6×indentation depth diameter [19]. the indentation covered space over the total investigated area (given by the grid dimensions 100 by 100 µm) was computed equal to 2.81% (ratio of efficiency). the second step (also displacement-driven with identical indentation depth level ii) was therefore designed for detailed investigation of selected position. the grid consisted of 25 by 25 indents separated by 2 µm and the efficiency equaled to 17.26%. the setting (especially location and size of the position) aimed for close evaluation of phases present in cement paste. 4. results interpretation the sem epma edx analysis of 41 material positions showed five main phases present in the investigated material, i.e. low-density and high-density calcium-silica-hydrate (ld c-s-h and hd c-s-h), calcium hydroxide (ch), non-hydrated clinker and aggregate – potassium and sodium feldspar and quartz (fig. 2). the phases average weight percentage of each element are summarized in tab. 1. moreover, 1.25 mm by 1.25 mm area, where level i indentation positions are expected, was sequentially figure 2. micrographs with identified material phases. scanned in bse mode with high resolution (5×5 sequence of 250 µm wide images). the reconstruction of this area enabled exact linkage between indentation data material structure. the histograms of nano-indentation results, the indentation modulus in particular, were normalized in respect to the 0.1 of minimum value, as captured for all studied levels in fig. 3. the preliminary study of the material (fig. 3a) proved to insufficiently detail mechanical properties of all phases. (a) . reduced modulus histogram – preliminary study. (b) . reduced modulus histogram – level i. (c) . reduced modulus histogram – level ii. figure 3. histograms of indentation modulus this can be observed in comparison with level i 57 v. hrbek, v. koudelková, z. prošek, p. tesárek acta polytechnica ctu proceedings element quartz potassium sodium ld c-s-h hd c-s-h ch clinker feldspar feldspar ca 6.65 37.37 45.37 60.05 50.18 si 46.74 29.39 27.24 6.41 12.83 0.35 12.42 al 10.09 11.06 1.42 1.33 0.50 mg 0.35 0.71 0.32 0.40 na 0.87 8.09 0.11 0.24 k 14.01 0.52 0.33 0.35 0.52 s 2.24 0.81 0.16 cl 0.36 fe 1.21 1.05 1.13 0.48 o 53.26 45.63 46.45 50.56 36.95 37.99 35.50 table 1. average weight percentage of elements present in phases. of the study (fig. 3b). the results, in this case, are more likely to depict the material structure as the occurrence of reduced modulus around 40 gpa is most frequent (see [23–26]). also, several peaks corresponding to number of phases identified by sem analysis can be found in the histogram. the level ii of the material was focused mostly on the investigation of cement paste, especially on csh and ch phase. again, fig. 3c confirms predominant representation of high-density calcium-silica-hydrate in the matrix. the overall results (fig. 4), as combination of both measured levels, gives detailed view of micromechanical performance of the material. the probability density function of measured indentation moduli was subsequently processed with spectral deconvolution in order to obtain gaussian distribution of expected phases. the mechanical properties of phases 1-5 in fig. 4 were thus determined as: phase 1 = 19.17 ± 3.54 gpa = ld c-s-h phase 2 = 39.75 ± 5.98 gpa = hd c-s-h phase 3 = 58.59 ± 6.34 gpa = calcium hydroxide phase 4 = 79.78 ± 3.11 gpa = presumably aggregate phase 5 = 118.35 ± 16.85 gpa = clinker figure 4. histogram of overall reduced modulus results. the results are consistent with previously published researches [23–32]. based on these results, maps of indentation moduli over the investigated positions (fig. 5 –6) were constructed with bordering conditions figure 5. contour map of preliminary results. figure 6. contour map of results on level ii. identical to limits in fig. 4. in addition, it was possible to link maps of indentation data to particular segments of sem bse micrographs (fig. 7). 58 vol. 13/2017 re-cycled concrete figure 7. contour maps of level i results linked to the sem bse micrograph. 5. conclusion the raw material (railway sleeper pb8) used after milling as recycled pulverized concrete in cementitious composites was in focus of this study. the micromechanical properties and material structure of this base material were investigated with combined methodology to predict performance when used as filler and cement replacement. the scanning electron microscopy methods (sem epma edx, bse) were used to determine chemical and structural composition of the material (unknown prior to the investigation). therefore, lowand high-density calcium-silica-hydrate (ld / hd c-s-h), calcium hydroxide (ch), three aggregate types and non-hydrated clinker were identified as main material phases. the nanoindentation technique over two main material levels enabled evaluation of the indentation modulus of these phases due to the spectral deconvolution of measured data. reverse reconstruction of indented areas with respect to deconvolution results let to interlink of indentation and bse diagrams. although the combined investigation provided sufficient results, more detail should be obtained in case of aggregate phases prior to modeling of the material. list of symbols f applied force [µn] hi indentation depth [nm] er reduced modulus [gpa] h hardness [gpa] hc contact depth [nm] acknowledgements the research is financially supported by czech technical university in prague – project sgs17/168/ohk1/3t/11 and by the czech science foundation research project 17-06771s. the support is gratefully appreciated. references [1] m. stibůrek. možnosti využití stavebních odpadů. beton tks 2:45–47, 2008. 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http://dx.doi.org/https://doi.org/10.1016/j.eml.2016.03.002 http://dx.doi.org/https://doi.org/10.1016/s0008-8846(03)00039-5 http://dx.doi.org/https://doi.org/10.1016/j.matchar.2007.05.018 http://dx.doi.org/https://doi.org/10.1016/j.cemconres.2008.09.002 http://dx.doi.org/https://doi.org/10.1016/j.cemconres.2010.09.001 http://dx.doi.org/https://doi.org/10.1007/s11340-010-9373-z http://dx.doi.org/https://doi.org/10.1016/s0008-8846(00)00505-6 acta polytechnica ctu proceedings 13:56–61, 2017 1 introduction 2 material 3 investigation procedure 3.1 scanning electron microscopy 3.2 quasi-static nano-indentation 4 results interpretation 5 conclusion list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.18.0048 acta polytechnica ctu proceedings 18:48–54, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app modelling and simulation of lattice structures using various material models for polymeric materials kerstin lehner∗, anna kalteis, zoltan major institute of polymer product engineering, johannes kepler university linz, altenbergerstraße 69, 4040 linz, austria ∗ corresponding author: kerstin.lehner@jku.at abstract. based on lightweight design concepts, lattices are increasingly considered as internal structures. this work deals with the simulation of periodically constructed lattices to characterize their behaviour under different loadings considering various material models. a thermo-mechanical analysis was done, which is resulting in negative clte-values (coefficient of linear thermal expansion). simulations with linear-elastic behaviour were evaluated regarding the tensile, compression and shear modulus and the poisson’s ratio. some of the investigated structures behave auxetic. beside the linear elastic behaviour, also the hyper-elastic and visco-elastic behaviour of some structures were investigated. furthermore, elasto-plastic simulations were performed where the applied loading was biaxial. as a result the initial yield surfaces were presented. the individual rves (representative volume elements) can be utilized for different areas of application dependent on the used materials. keywords: periodic lattice structures, material models, auxetic behaviour, negative clte. 1. introduction based on lightweight design concepts, lattices are increasingly considered for additively manufactured components as internal structures in spite of classical bulk solid volumes. the structured construction of cellular materials has extremely advantageous properties for the lightweight design represented by a low material density combined with high stiffness and strength as well as an excellent damping behaviour [1]. dependent on the desired applications, various materials possessing different behavioural mechanisms are utilized for the design of components. the aim of the present work was to characterize and compare various modelled regular open cellular materials defined as lattice structures with applied periodic boundary conditions regarding different material models under various loading conditions. fig.1 depicts the task and procedure of the work performed. the results of the structures which are presented in this paper are highlighted. these results are chosen because they are the most informative ones. linear-elastic material behaviour is the simplest assumption of elasticity and occurs for small deformations [2]. thermoplastic elastomers behave hyper-elastically, which is characterized by a non-linear deformation behaviour and a unique stored strain energy density function depending on the current strain [3]. visco-elastic materials behave both viscous and elastic and show time-, temperatureand frequency-dependent characteristics [4]. elasto-plasticity describes polymeric materials, in which plastic deformation occurs when loaded beyond its limit of linear-elastic behaviour. in the next section the simulation approaches and results are presented and in the last section a conclusion for various lattice structures and different material models is drawn. 2. methodology in the following the methodology is described how the lattice structures are modelled and simulated. the structures are modelled in the software nx, which was developed by the company siemens. it is an interactive cad/cam/ cae-system. fig.2 shows the procedure of the modelling of the various structures represented by an octet structure. a matrix block is modelled by extrusion of a square with the dimensions 50 x 50 mm2. within this assembly space sketches are drawn along the desired geometry of the trusses. these are fully parametrised, which is when the size of the representative volume element (rve) is changed the inclusion size is automatically adapted. at these curves filled tubes are positioned, which are connected to each other. boolean operations were applied to first intersect the protruded inclusion parts from the rve and then subtract the inclusion from the matrix to get two individual parts. with these structures finite element simulations are performed in the software abaqus cae (simulia, dassault systemes). in order to model an rve in such a way that there is no influence of edge effects, which means that the rve has to be modelled as if being part of the continuous bulk material, commonly so called pbcs are used. using pbcs the single rve behaves as if being part of a continuum material [5]. therefore, perodic boundary conditions (pbcs) are implemented in abaqus cae by a python script, which was developed at the institute of polymer product engineering (ippe) by 48 http://dx.doi.org/10.14311/app.2018.18.0048 http://ojs.cvut.cz/ojs/index.php/app vol. 18/2018 modelling and simulation of lattice structures figure 1. task and procedure of the present work. dipl.-ing. dr. matei-constantin miron and dipl.ing. emil jacob pitz. the specific procedure and mechanism of this script can be found in [6]. 3. thermo-mechanical analysis thermal expansion is the tendency of matter to change in shape, area and volume in response to a change in temperature [7]. in the case of linear observation, the one-dimensional change in length is discussed. the strain in the direction of the length, which is equally to the relative expansion, divided by the change in temperature is called the material’s coefficient of linear thermal expansion (clte) and varies with temperature for most materials. the mathematical formula for the clte α is given in eq.1. ∆l is the change in length (∆l = l1 − l0), l0 is the initial length of the specimen and ∆t = t1 −t0 represents the temperature change. α = ∆l l0 · 1 ∆t (1) for clte-simulations predefined fields with a constant temperature rise of 10 k throughout the whole region are implemented. the reference points for the three directions in space, which are created by a script for periodic boundary conditions and mark the centres of the adjacent unit cells, are set free to permit expansion in all directions. the partitioning of structure 9 for the two-phase system is given in fig.3 with the outer part coloured grey and the inner part coloured purple. two different variations of structure 9 are investigated: structure 9a (tube diameter of 5 mm) and structure 9d (tube diameter of 2.5 mm). the used material data for the thermo-mechanical simulations are given in tab.1. material 1 has a lower e-modulus, a lower α and a higher ν than material 2. the resulting von mises stresses can be seen in fig.4 (top) for a material composition where the inner square is made material 1 material 2 e,mpa 1000 10000 ν, 0.45 0.2 α, 1 k 1 e-6 1 e-5 table 1. material data for clte-simulations [8]. of a stiffer material with a higher coefficient of linear thermal expansion (material 2) than the outer structural part. the maximum stresses occur at the connection points because of the hindered expansion of the inner square due to the lower expansion of the outer parts. fig.4 (bottom) shows the reversed construction where the inner square consists of material 1. greater stresses occur in the outer geometry part. when using material 1 for the outer structural part the cltes yield in negative values for the two variations of structure 9 and therefore contraction in z-direction (fig. 5, 1. and 2.). this effect is traced to a mechanical contraction normal to the plane in which the inner square consisting of material 2 lies resulting from the higher expansion of the inner material. this mechanical contraction overcomes the expansion due to the temperature increase and therefore yields to a negative clte in z-direction for this structure. compared to the literature results of [8], this lattice structure reveals nearly the same value of the thermal expansion in z-direction. when the outer structural part is made out of material 2 and the inner square consists of material 1 both investigated structures expand when increasing the temperature (fig. 5, 3. and 4.). for this material composition the highest clte-values are received in z-direction. 4. linear-elasticity linear-elasticity describes the material behaviour at small deformations [2]. most elastic materials exhibit 49 k. lehner, a. kalteis, z. major acta polytechnica ctu proceedings figure 2. modelling procedure represented by an octet structure. figure 3. partitioning of structure 9. a linear-elastic model, which is time-independent, and can be described by a linear relation between the stress σ and strain �. linear-elastic behaviour for the lattice structures is characterized by the apparent lattice stiffness matrix wherefore displacement-controlled simulations of uniaxial tension (et ) and shearing (g) were performed in all directions to determine the entries of the stiffness matrix. for further investigations of the linear behaviour also simulations of uniaxial compression (ec) were done. the following input parameters are used for the linear-elastic simulations: e = 72000 mpa, ν = 0.2. in fig.6 it can be seen that structure 6 shows the stiffest behaviour of all investigated structures in all directions. structure 3 has the lowest stiffness values related to tension and compression, but this rve shows a good applicability for purposes where shear loads are present. when comparing the different variations of structure 9 it can be seen that structure 9a (tube diameter of 5 mm) results in the highest stiffness values whilst structure 9c (hollow tubes with a shell thickness of 0.5 mm) figure 4. von mises stresses of a two-phase system caused by temperature increase. and structure 9d (tube diameter of 2.5 mm) exhibit lower values, but also have the maxima in z-direction. special characteristics of the poisson’s ratio can be observed for structure 5 (fig.7 (top)), structure 8 (fig.7 (bottom left)) and structure 10 (fig.7 (bottom right)). in fig.7 these structures are illustrated and the von mises stress is depicted for different loading directions. structure 5 expands in y-direction when pulling in x-direction and vice versa, which is described by negative values for ν12 and ν21 (auxetic behaviour). with tension in z-direction this structure bears very high strains in xand y-direction. structure 8 shows an overall auxetic behaviour. due to the twisting mo50 vol. 18/2018 modelling and simulation of lattice structures figure 5. von mises stresses of a two-phase system caused by temperature increase. figure 6. comparison of the different structures regarding tensile (et 1), compression (ec1) and shear (g12) modulus. tion, structure 10 also behaves auxetic and therefore exhibits negative values for the poisson’s ratios. 5. hyper-elasticity hyper-elasticity represents a non-linear, isotropic and time-independent form of the elasticity. the relationship between an applied stress and the strain is linear for very small strains, but will reach a plateau in the stress-strain curve at a certain point [3]. the strain energy density function w is evaluated using the displacements resulting from the forcecontrolled simulations according to the law of neohookean solids, which is presented in eq.2 where c10 is a material constant and i1 is the first invariant of the stress tensor. the hyper-elastic material model (neo-hooke) was generated from experimental stress-strain-data resulted from uniaxial tensile tests of the bulk materials ninjaflex® (e = 6.5 mpa, ν = 0.495) and tpu3 (e = 9.5 mpa, ν = 0.495). the material constant c10 for the neo-hookean model figure 7. illustration of the auxetic behaviour of structure 5, 8 and 10. figure 8. comparison of w for the different lattice structures and for the two materials. has the value c10 = 0.78057 nmm2 for ninjaflex ® and c10 = 1.56371 nmm2 for tpu3. wn h = c10 · (i1 − 3) (2) a bar chart of the w-values of the various relevant structures is given in fig.8. it can be seen that structure 10 reveals the highest values for w, which means that the applied force can be stored in the material as an energy to change the shape of the structure. therefore, this structure is the best fitting structure of all investigated rves for hyper-elastic characteristics. 6. visco-elasticity visco-elastic materials combine the characteristics of solids and liquids, and thus behave in part elastically and in part plastically. the effect of visco-elasticity is time-, temperatureand frequency-dependent. to determine the visco-elastic material behaviour two strainrates were evaluated. the simulations were performed 51 k. lehner, a. kalteis, z. major acta polytechnica ctu proceedings dynamic, implicit to allow strain-rate dependent calculations. the applied loadings were 0.5 mm s and 5 mm s for uniaxial tension. for the visco-elastic simulations experimental data from a dynamic-mechanical thermal analysis (dmta) were adapted into master curves with the help of the program viscoshift (ingenieurbuero herdy) and the prony parameters (tab.2) were calculated in the program viscodata (ingenieurbuero herdy). the directly calculated parameters ei,p rony, ki,p rony and τi,p rony are representing the young’s modulus, the compression modulus and the relaxation time, while the shear modulus gi,p rony has to be calculated by gi,p rony = ei,p rony 2 · (1 + ν) when assuming linear visco-elastic behaviour. the details of the exact implementation of the visco-elastic material behaviour into abaqus can be found in [9]. in order to be able to compare the results, these simulations were performed for structure 10 as well as for a multipurpose tensile specimen according to iso 527-1. in fig.9 it can be seen that an increased strain-rate leads to a stiffer behaviour for the investigated material independent of the structure. when comparing the two specimens the results show lower stress values for structure 10 (fig.9 (top)) than for the multi-purpose tensile specimen (fig.9 (bottom)), which represents the bulk material. this effect is traced to the low relative density of the lattice structure. 7. elasto-plasticity in real media, deformation is linearly elastic only up to a certain limit. if this limit is exceeded and the material is therefore loaded beyond its yield stress, in ductile materials plastic deformation occurs [8]. after reaching the yield point, which is the point on the stress-strain curve that indicates the limit of elastic behaviour and the beginning of plastic behaviour, the stressed body does not return to its original form when it gets unloaded. for the characterization of elastoplastic material behaviour the initial yield surfaces were determined. therefore simulations of biaxial loading conditions and coupled tension/compressionsimulations were necessary. the beginning of plasticity is assumed where a deviation of the experimental curve from linear behaviour can be detected to represent the real curve progression. when considering the physical yield point at approximately 32 mpa the simulation model would be linear up to this point and would therefore not be suitable as approximation for the elasto-plastic behaviour. hence, the plastic strain and the yield stress for pp(h) (e = 1990 mpa, ν = 0.42), which represent elasto-plastic material behaviour, are shown in fig.10 (yield point) and implemented in abaqus using mises-plasticity without consideration of triaxiality. in fig.11 the rves are pictured on the left side showing the locations of the beginning of yielding in red for biaxial tension in yand z-direction. the blue colour indicates elastic behaviour. on the right side the yield points for figure 9. illustration of visco-elastic behaviour of str. 10 (top) and iso-mps (bottom). the different load cases are connected with straight lines. these lines mark the border between elastic and plastic behaviour. if the stresses exceed this border, plasticity is present. additionally, the initial yield surfaces resulting from the yield criteria of von mises and tresca are plotted. the tresca criterion (eq.3) is marked with a dotted line and the von mises criterion (eq.4) is drawn as a dashed line. σt = max { 1 2 |σ1 −σ2|, 1 2 |σ2 −σ3|, 1 2 |σ3 −σ1| } (3) σ2v = 1 2 · [(σ11 −σ22)2 + (σ22 −σ33)2 + (σ33 −σ11)2 + 6 · (σ223 + σ 2 31 + σ 2 12)] (4) the initial yielding of structure 3 (fig.11(top)) takes place in its centre and therefore possesses lower yield stresses than structure 4 (fig.11(top centre)), the centre of which is separated into more connection points. structure 8 (fig.11(bottom)) shows the highest stress values for the transition between elasticity and plasticity for biaxial loading with tension in one 52 vol. 18/2018 modelling and simulation of lattice structures ei,p rony,− gi,p rony,− ki,p rony,− τi,p rony,s 2.38 · 10−18 7.96 · 10−19 7.93 · 10−17 1.29 · 10−03 1.40 · 10−18 4.68 · 10−19 4.67 · 10−17 2.98 · 10−02 6.44 · 10−18 2.15 · 10−18 2.15 · 10−16 6.85 · 10−01 1.27 · 10−17 4.24 · 10−18 4.22 · 10−16 1.58 · 10+01 8.33 · 10−20 2.79 · 10−20 2.78 · 10−18 3.63 · 10+02 2.49 · 10−17 8.33 · 10−18 8.30 · 10−16 8.36 · 10+03 6.92 · 10−17 2.32 · 10−17 2.31 · 10−15 1.92 · 10+05 3.97 · 10−16 1.33 · 10−16 1.32 · 10−14 4.43 · 10+06 2.11 · 10−02 7.04 · 10−03 7.02 · 10−01 1.02 · 10+08 table 2. prony parameters: ninjaflex®. figure 10. experimental data for elasto-plastic material model: pp(h). direction and simultaneous compression in the other direction, whereas all non-auxetic structures have the highest yield stresses for loadings with the same sign. for structure 4 (fig.11(top centre)) the tresca yield criterion is very similar to the resulting yield surface and for structure 6 (fig.11(bottom centre)) the von mises criterion would be a good approximation. 8. conclusion performing a thermo-mechanical analysis simulations with partitioned inclusion geometries, which were applied with two different materials with various expansion behaviour, were conducted. the findings of [8] that a specific structural composition leads to a negative coefficient of linear thermal expansion could be validated for two-phase systems. regarding the parameters of the apparent lattice stiffness matrix, structure 6 was presented as the stiffest of all investigated structures in all directions using linear-elastic material behaviour. the lowest stiffness values related to tension and compression were detected for structure 3. negative values for the poisson’s ratio, which characterizes auxetic behaviour, could be observed for structure 5, structure 8 and structure 10. after performing force-controlled simulations with hyperelastic material behaviour the strain energy density function was calculated according to the law of neohookean solids. structure 10 revealed the highest values for w, which made this rve most suitable for large-strain non-linear elastic material behaviour. figure 11. initial yield surface: str. 3, 4, 6, 8. 53 k. lehner, a. kalteis, z. major acta polytechnica ctu proceedings visco-elastic simulations were performed using prony parameters of a generalized maxwell model to investigate the strain rate dependency of structure 10 compared to a multi-purpose tensile specimen. whilst an increased strain-rate led to a stiffer behaviour for the investigated material independent of the structure, structure 10 generally resulted in lower stress values than the bulk material. in order to represent the results of elasto-plastic material behaviour, initial yield surfaces were determined by applying biaxial loading conditions. in contrast to all investigated geometries, structure 2 and structure 3 exhibited the lowest yield stresses while structure 5 and structure 6 could withstand the highest stresses before yielding. all in all it was shown that the various lattice structures show different thermal and mechanical behaviour and therefore can be utilized for various purposes in different areas of application dependent on the used materials. references [1] k. drechsler, i. karb, r. kehrle, v. witzel. faserverbundstrukturen, leichtbau in natur und technik. themenheft forschung leichtbau pp. 14–21, 2007. [2] p. kelly. solid mechanics part i: an introduction to solid mechanics, chap. 6, pp. 145–153. 2013. [3] m. boyce, e. arruda. constitutive models of rubber elasticity: a review. rubber chemistry and technology 73(3):504–523, 2000. [4] m. blanter, i. golovin, h. neuhäuser, h. sinning. internal friction in metallic materials, vol. 90. springer-verlag berlin heidelberg, 2007. [5] v. d. nguyen. computational homogenization of cellular materials capturing micro-buckling, macro-localization and size effects. ph.d. thesis, universite de liege computational and multiscale mechanics of materials. 2014. [6] e. j. pitz. multiscale progressive damage modelling in continuously reinforced composite structures. master thesis, johannes kepler university linz, 2017. [7] p. a. tipler, g. mosca. physics for scientists and engineers, vol. 1. macmillan education, 2007. [8] j. aboudi, r. gilat. micromechanical analysis of lattice blocks. international journal of solids and structures 42(15):4372–4392, 2005. [9] m. machado, u. cakmak, i. kallai, z. major. thermomechanical viscoelastic analysis of woven-reinforced thermoplastic-matrix composites. composite structures 157:256–264, 2016. 54 acta polytechnica ctu proceedings 18:48–54, 2018 1 introduction 2 methodology 3 thermo-mechanical analysis 4 linear-elasticity 5 hyper-elasticity 6 visco-elasticity 7 elasto-plasticity 8 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2018.20.0051 acta polytechnica ctu proceedings 20:51–55, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of total pressure fluctuations in steam turbine stages with different degree of reaction tomáš jelíneka, ∗, martin němeca, petr milčákb a czech aerospace research centre, beranových 130, prague 9, czech republic b doosan škoda power, tylova 1/57, pilsen, czech republic ∗ corresponding author: jelinek@vzlu.cz abstract. the contribution deals with analysis of unsteady pressure measurement in one and half axial turbine stages. the measurement was performed in vzlu air test turbine. two stages geometries with different degree of reaction were tested in this work. comparison of unsteady flow field behaviours is discussed. keywords: turbine stage, stator-rotor interaction, driving forces. 1. introduction a multistage arrangement design of turbines is more and more complex and it is necessary to combine aerodynamic and strength aspects. this evolution of design procedures is followed also in a long-term cooperation of vzlu with the leading producer of steam turbines doosan škoda power (dspw). blades excitation is strongly affected by axial gap, blade count ratio and overall design of the stage. combined effect of the axial gap and blade count ratio was deeply studied by chen et al. in [1]. their results show resonance effect of stator and rotor row for huge number of combinations of blade count ratios and axial gaps. the resonance effect is caused by interaction of the rotor potential field with the stator outlet flow field. moreover, in the case of a multistage turbine, the interstage interaction and clocking effect influences the flow field development what presented reinmöller at al. in [2] and porreca at al. in [3]. a sensitivity of stage geometry on the excitation is presented by gao et al. in [4], where decrease of stimulus was achieved by using a negative bowed stator geometry. many other works are focused on optimization of the stage geometry. such studies are performed with help of numerical simulations and usually it is necessary to use scaling a number of blades to reduce computational costs. this could generate a fairly large discrepancy as showed clark in [5]. this work is unique because of two degrees of reaction and various stator aspect ratios were experimentally investigated on one and half stage configuration. experimentally obtained unsteady flow field parameters were analyzed and comparison of the stimulus behaviours of novel steam turbine stage design was performed. driving forces were not determined, but only pressure fluctuations were computed from unsteady flow field parameters measurements. this approach is sufficient for the comparison of various cases. a reaction stage and a stage with increased reaction figure 1. the one and half stage cofiguration of the test rubine. (marked as reac and ir, respectively) were studied in nominal and off-design regimes. the experiments caught unsteady flow field behaviours at the outlet of the rotor and also at the outlet of the second stator. an influence of degree of reaction in combination with stator aspect ratio and rotor rotational speed is studied in this work. 2. test turbine configuration the experimental work was performed in vzlu in the axial turbine stage test rig. the rig was adapted to the one and half turbine stage to study the multistage flow structures [6]. the one and half stage facility is depicted in fig. 1. the configurations of tested stages are characterized in tab. 1. an axial gap ratio (gr) and a vane/blade count ratio (bcr) are important parameters in stimulus point of view, therefore these parameters are also included in this table. detailed flow field investigation was performed by a miniature five-hole pressure probes (tip diameter 1.2 mm) and with a fast response pressure probes (frpp) developed in vzlu [7]. 51 http://dx.doi.org/10.14311/app.2018.20.0051 http://ojs.cvut.cz/ojs/index.php/app t. jelínek, m. němec, p. milčák acta polytechnica ctu proceedings reac v2-1 & v2-2 ir v1-2 & v1-2 stator aspect ratio 1st / 2nd 2.0 / 1.1 1.0 / 1.0 rotor aspect ratio 1.71 1.94 stage isentropic mach number 0.30 0.40 mid-span reaction 0.43 0.12 optimum velocity ratio u/cis 0.65 0.55 midspan gap ratio plane 1 / 2 0.23 / 0.67 0.11 / 0.63 blade count ratio plane 1 / 2 0.61 / 1.09 0.53 / 1.88 table 1. the turbine stage geometry and operational regime parameters. (a) . reac stage. (b) . ir stage. figure 2. the first rotor excitation for two stator aspect ratios. 3. the first stator excitation the first stator is specified by uniform inlet flow in this test turbine configuration, so the excitation of vanes is caused only due to rotor potential field. it is shown in [1], that the worst case occurs for bcr approximately 0.7, where the resonance appears. the bcr is sufficiently far from the resonance in this case. the excitation by the rotor potential field is captured in the case of the ir stage also by the fast response pressure probe placed in the plane 1, see fig. 1. measured amplitude is very low what is consistent with bcr value 0.53. radial distribution of the absolute total pressure amplitude is shown in fig. 4b. the absolute value of the amplitude is approximately four times lower compared to the rotor excitation, that is in agreement with [1]. influence of the rotational speed on the excitation of the first stator is shown in fig. 6a. 4. the first rotor excitation the first rotor excitation is mainly caused by passing the first stator wakes. the excitation of four cases is shown in following graphs fig. 2a and fig. 2b. these graphs show comparison of two stator families and two aspect ratios for each family. high aspect ratio stator geometry is characterised by a stronger decrease of the amplitude with increasing degree of harmonic component. low aspect ratio stator has a very similar excitation for three harmonics in the midspan region. the highest excitation is captured in secondary flow regions. 5. the second stator excitation the excitation of the second stator is caused by rotor wakes at the inlet of the stator. the relative pressure fluctuations are the highest in turbine stage (shown in fig. 5a and fig. 5b). nevertheless, the second stator excitation in absolute values point of view is approximately three times lower compared to the rotor excitation radial distribution is shown in fig. 4a and fig. 4b amplitude scale is kept the same in both graphs. the excitation of the second stator at the outlet caused by the first rotor wakes is slightly lower compared to the inlet excitation. as the second rotor is not present in this configuration, there is no excitation by second rotor potential field, which could cause additional stator excitation. the second stator excitation is moreover dependent on the clocking with the first stator. clocking effects are discussed in [8] for the case of the reac stage. 6. the second rotor excitation although the second rotor was not present in the test turbine, the excitation of the second stage rotor can be obtained from the second stator time averaged outlet flow field. radial distribution of the rotor excitation is shown in fig. 3a (relative values) and in fig. 3b (absolute values). both second stator geometries had 52 vol. 20/2018 comparison of total pressure fluctuations in steam turbine stages (a) . relative total pressure amplitudes. (b) . absolute total pressure amplitudes. figure 3. the second rotor excitation reac and ir stage. (a) . reac stage. (b) . ir stage. figure 4. absolute total pressure amplitudes. (a) . reac stage. (b) . ir stage. figure 5. relative total pressure amplitudes. 53 t. jelínek, m. němec, p. milčák acta polytechnica ctu proceedings (a) . the first stator of ir stage. (b) . the second stator of ir and reac stage. figure 6. influence of rotational speed on stators excitation. (a) . reac stage. (b) . ir stage. figure 7. the second rotor excitation influence of rotational speed. low aspect ratio, so a decrease of the amplitude with increase of harmonic component was less significant. 7. the rotor rotational speed influence the rotational speed strongly influences the flow filed in the turbine stage. as a blade incidence angle is changed, separation regions may occur, radial distribution of velocity is affected and thus unsteady flow field behaviours are changed. the potential field of the rotor is also slightly affected by the rotational speed, so the excitation of the first stator is also changed, as it is shown in fig. 6a. as it was said, the first rotor is excited due to its interaction with first stator wakes. the most important for the rotor excitation is time averaged distribution of the pressure at the stator outlet. the time averaged first stator pressure distribution is not seriously affected by the rotational speed, thus the excitation of the rotor is also not changed. the second stator excitation by the rotor outlet flow field is shown in fig. 6b. the variation of the excitation in radial distribution is clearly visible in this figure, namely for the ir stage. a lower expansion in the case of the ir stage causes a higher influences of the rotational speed on the second stator excitation. the highest excitation of the second stator is captured for nominal regimes. the second rotor excitation radial distribution with the rotational speed influence is shown in fig. 7a and fig. 7b for both stages. the radial distribution of the excitation is influenced more in the case of reac stage. extremely overloaded stage has a region of high excitation in the midspan. this effect is not as strong in the case of the ir stage, which is specified by more intensive expansion in the stator. 8. conclusion the excitation of one and half stages was studied for two degrees of reaction in wide range of regimes. the comparison of excitation results of stages with two degrees of reaction shown excitation insensitivity in the relative point of view. absolute values of excitation are of course higher in the case of lower reaction stage, because of higher dynamic pressure, which is used for the normalization. the first and the second rotor relative excitation stays nearly the same. this statement is valid for average excitation, because distribution of the excitation along blade height is strongly affected by the second stator secondary flow development. the second stator is strongly affected by the interaction with 54 vol. 20/2018 comparison of total pressure fluctuations in steam turbine stages rotor tip leakage flow field. in the case of off-design regimes, the secondary flow development is affected by a change of the second stator incidence angle. the second rotor absolute amplitudes are slightly higher compared to the first rotor. this is caused by a higher outlet mach number of the second stator compared to the first one. stator aspect ratio can influence damping of the excitation of higher harmonic components. it was tested in the case of the first rotor. excitation of the stator by the rotor potential field was also discussed and tested in the case of lower reaction stage. the results show that the excitation is the weakest in the nominal regime and grows with deviation from the nominal regime. the excitation caused by the potential field is generally weak for the cases where resonance does not occur. list of symbols cis isentropic stage velocity [m/s] l nondimensional blade height [1] q dynamic pressure [pa] u rotor circumferential velocity [m/s] ar aspect ratio bcr blade count ratio dsp w doosan škoda power f rp p fast response pressure probe gr axial gap ratio h harmonic component ir increased reaction stage reac reaction stage rp m nominal rotor rotational speed acknowledgements the work was done with financial support of the technology agency of the czech republic on the project ta04020129 "investigation of flows in a multi-stage axial turbine stages arrangement". references [1] t. chen, k. patel, p. millington. combined effects of both axial gap and blade count ratio on the unsteady forces of a steam turbine stage. asme paper no. gt2012-68874. [2] u. reinmöller, b. stephan, s. schmidt, n. r. clocking effects in a 1.5 stage axial turbine steady and unsteady experimental investigations supported by numerical simulations. asme paper no. 2001-gt-0304. [3] l. porreca, a. i. kalfas, r. s. abhari. interstage flow interactions and loss generation in a two-stage shrouded axial turbine. journal of turbomachinery 131(january), 2009. [4] k. gao, t. lin, z. xie, d. zhang. investigation on unsteady flow of a turbine stage with negative bowed stator. international journal of energy and power engineering 11(4), 2017. [5] j. p. clark, g. m. stetson, s. s. magge, et al. the effect of airfoil scaling on the predicted unsteady loading on the blade of a 1 and 1/2 stage transonic turbine and a comparison with experimental results. asme paper no. 2000-gt-0446. [6] m. nemec, t. jelinek, p. milcak. an investigation of a flow field in a one and half axial turbine stage. aip conference proceedings 1889(020025), 2017. [7] m. nemec. development of fast response pressure probe for turbomachinery applications (in czech). pse 2009, pilsen, 2009. [8] m. nemec, t. jelinek, p. milcak. clocking of stators in one and half stage of axial steam turbine. epj web of conferences 180(02071), 2018. 55 acta polytechnica ctu proceedings 20:51–55, 2018 1 introduction 2 test turbine configuration 3 the first stator excitation 4 the first rotor excitation 5 the second stator excitation 6 the second rotor excitation 7 the rotor rotational speed influence 8 conclusion list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.9.0032 acta polytechnica ctu proceedings 9:32–38, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app real structure and residual stresses in advanced welds determined by x-ray and neutron diffraction karel trojana, ∗, charles hervochesb, kamil kolaříka, nikolaj ganeva, pavol mikulab, jiří čapeka a ctu in prague, department of solid state engineering, faculty of nuclear sciences and physical engineering, trojanova 13, 120 00 prague 2, czech republic b nuclear physics institute, ascr, v.v.i., department of neutron physics, 25068 řež, czech republic ∗ corresponding author: karel.trojan@fjfi.cvut.cz abstract. the paper outlines the capability of x-ray diffraction (xrd) for evaluation of real structure changes and residual stresses (rs) on cross-section of advanced thick welds due to the welding of ferromagnetic plates. the results of neutron diffraction describe a three-dimensional state of rs and also verify previous assumptions of rs redistribution as a result of the surface preparation for determination 2d maps measured by xrd. keywords: laser and mag welding, residual stresses, x-ray diffraction, neutron diffraction. 1. introduction high local residual stresses existing in components could have vital influence on its properties. when a welded part with high local rs is machined the significant distortion may occur, due to the disruption of the equilibrium state of rs. on the other hand high tensile rs due to welding have a strong negative effect on the strength properties, especially under fatigue loading. therefore, the determination of the rs distribution in complex welded components is accentuated. recent practical approaches distinguish between local and global welding rs [1]. the local rs are resulting from the local heating and cooling processes in the weld metal and adjacent heat affected zone (haz). global rs exist in the entire component and they are resulting from shrinkage processes which occur in the whole component [1]. these global rs are affected by the stiffness of the component and the local shrinkage processes in the direction of black arrows in fig. 1. tensile rs around a single pass weld in a plane sample are expected due to the restraint which is generated by the cool adjacent zones in the base material of the plate [1]. therefore, the main direction of tensile rs due to shrinkage is applied along the weld. in the perpendicular direction, the tensile rs should reach lower values according to this model. for a complete description of the austenite phase, the transformation associated with a change in volume must be taken into consideration as a significant source of rs especially in high strength steels. the local compressive rs, which arise in the weld zone as a consequence of the restrained volume expansion during the transformation of austenite in martensite, bainite or ferrite, are superimposed to the global tensile rs figure 1. shrinkage tendency and constraint conditions in a single pass weld (schematically) [1]. explained according to the model described above (fig. 1) [1]. the increase in the cooling rate of the welded plates will lead to quenching and formation of local compressive rs. laser welding supplies into the weld lower heat input so that cooling of the weld is faster compared to the mag technology. the resulting distribution of residual stresses is a combined effect of hindered shrinkage and phase transformations. developed laser welding methods using high power diode lasers (hpdl) took over the capability to fill groove with cold or hot wire metal from metal active gas (mag) welding and thus to change mechanical properties of welds reducing their hardness due to the quenching [2]. profitable changes of real crystallographic structure and rs in comparison with conventional laser welds improve the results during impact and tensile test and mainly enhance fatigue life as a result of a favourable distribution of rs and its relaxation during welding in the weld zone and haz [3]. 32 http://dx.doi.org/10.14311/app.2017.9.0032 http://ojs.cvut.cz/ojs/index.php/app vol. 9/2017 real structure and residual stresses in advanced welds to verify the results obtained by xrd, where the samples were cut in two pieces, it is preferable to use the neutron diffraction. it has distinguished benefits such as the unique deep-penetration, threedimensional mapping capability, and volume-averaged bulk measurements characteristic of the scattering neutron beam for steel sample thickness up to 50 mm without cutting [4]. 2. experimental methods the analysed samples were prepared by a hpdl laser with cold wire and mag welding (metal active gas (mag) welding subtype of gas metal arc welding) from two 20 mm thick sheets made of s355j2 steel. the xrd measurements were performed by proto ixrd combo diffractometer with ω-goniometer and {211} diffraction line of α-fe was measured by crkα radiation. deformation of interplanar distances of variously rotated planes was converted to rs using the generalized hooke’s law according to the sin2ψ method. the corresponding neutron measurements were carried out on the diffractometer spn-100 of the nuclear physics institute which is installed at the channel hc-4 of the research reactor lvr-15 and operates at the neutron wavelength of 0.235 nm. the steca software [5] was used to extract and analyse diffraction patterns from the recorded area detector data. the diffraction peak {110} of α-fe was measured. the strain was calculated using the hooke’s law based on the angular deviation of the diffraction profile position from the value related to the stress-free sample. full widths of the measured diffraction lines at half of the maximum (fwhm ) were also evaluated for both methods. fwhm parameter depends on microstrains, crystallite size, density of dislocations and weight proportion of the hard phases. sample orientations during xrd and neutron diffraction are shown in figs. 2a and 2b. the direction of the arrows always indicates the direction of the measured rs. the rotation of measured planes during sin2ψ measurements is also shown in fig. 2a for xrd. equivalent direction are therefore l’ for xrd and t for neutron diffracton. the measurements were performed in three lines perpendicular to the welds. one line passes through the centre of the weld and the other two are located three millimetres below each surface, see red lines in fig. 2c. upper side of the plate is the site which was firstly welded. to determine the real structure and rs by xrd on the cross section, it was necessary to cut the plate in two parts (solid lines and dashed lines in fig. 2c) and the affected surface layer electrochemically etched. 3. experiment gradients of rs measured by xrd in two directions on the surface of welded plates with a depth of penetration of the used radiation into the sample approx. (a) . sample orientations during xrd. (b) . sample orientations during neutron diffraction. (c) . measured lines during neutron diffraction. figure 2. sample orientations during experimental methods. 4 µm are plotted in fig. 3. dependences of threedimensional state of rs determined by neutron diffraction technique related to the middle line and two lines with depth of three millimetres below both surfaces of laser and mag samples can be seen in figs. 4 and 5. the comparison of rs obtained by xrd in direction l’ and neutron diffraction in direction t is plotted in figs. 6 and 7. the comparison of fwhm obtained 33 k. trojan, ch. hervoches, k. kolařík et al. acta polytechnica ctu proceedings figure 3. surface residual stresses measured by xrd. by xrd in direction l’ and neutron diffraction in direction t is shown in figs. 8 and 9. the parameter x is the perpendicular distance from the axis of the weld in all graphs. the most significant tensile rs on the surface are in the direction s according to the predictions (see fig. 1 and 3) for both samples. tensile rs of the laser weld in the direction s exceed the yield strength of the used steel and for upper side even the tensile strength. this would suggest that the rapid cooling of laser weld created non-equilibrium hard phases with higher strength [6]. in l direction differently wide haz could be well observed. according to figs. 4 and 5 tensile rs within all the three measured lines obtained by neutron diffraction have the maximum in the longitudinal direction for both samples. rs in the laser weld at middle line exceeding 300 mpa and for mag in bottom line only 230 mpa. in the middle of the laser and mag weld, the distribution of rs in both remaining directions n and t is the same. shrinkage and other effects caused that rs were created in both directions homogeneously. in the middle of the laser weld according to figs. 6 and 7, rs obtained by neutron diffraction in the perpendicular direction to the weld reported more tension character than by xrd. this is most probably due to a significant contraction in the direction parallel to the weld. on the other hand in the case of the mag weld, there are compressive stresses. this could suggest a predominance of phase transformation during the formation of rs rather than shrinkage. for both welds, rs obtained by xrd reported higher compressive results than obtained by neutron diffraction. the only exception is the upper line for the mag weld, see fig 7. laser weld (according to the fig. 4) has a higher tensile residual stresses along the weld, therefore, redistribution after cutting the sample probably produced the greatest compressive stresses measured by xrd. after cutting the plate, the weld probably slightly dropped, and thus it caused compressive rs in the plane perpendicular to the weld. the comparison of the parameter fwhm using both methods exhibits in nearly all cases the same pattern, see figs. 8 and 9. change in the parameter fwhm for mag sample, which can be correlated with a width of haz, is recorded at a greater distance than for the laser weld. fwhm is in all cases for both methods higher for the laser sample, this would indicate a higher microstrain, density of dislocations and a greater amount of hard phases. 4. conclusions it was found from the residual stresses measured by neutron diffraction in three different sample orientations and in three different depths under the surface that the greatest gradient of residual stresses is in the direction parallel to the weld. according to the literature study and our results, greater tensile residual stresses exhibits the laser weld in almost all measured lines. furthermore, the results describing haz width (changes in values of rs and fwhm ) are very similar for both analytical methods (xrd and neutron diffraction). both diffraction methods are able to complement each other. furthermore, the resulting residual stresses gradients obtained using both methods with a different approach are in certain correlation. each method (xrd and neutron diffraction) describes the outcome of residual stresses in different context. the redistribution of rs during the cutting has not yet been accurately described, so further measurements and verification using mathematical modelling will be needed. we are currently performing the validation of measurement on the second mag sheet. acknowledgements measurements were carried out at the canam infrastructure of the npi ascr řež supported through mšmt project no. lm2011019 and czech science foundation gačr through the project no. 14-36566g entitled as multidisciplinary research centre for advanced materials. this work was supported by the student grant competition ctu in prague grant no. sgs16/245/ohk4/3t/14. the authors thank to matex pm for supplying the laser weld. 34 vol. 9/2017 real structure and residual stresses in advanced welds (a) . laser — upper line. (b) . laser — middle line. (c) . laser — bottom line. figure 4. rs for laser sample along three measured lines obtained by neutron diffraction. (a) . mag — upper line. (b) . mag — middle line. (c) . mag — bottom line. figure 5. rs for mag sample along three measured lines obtained by neutron diffraction. 35 k. trojan, ch. hervoches, k. kolařík et al. acta polytechnica ctu proceedings (a) . laser — upper line. (b) . laser — middle line. (c) . laser — bottom line. figure 6. comparison of rs obtained by xrd in direction l’ and neutron diffraction in direction t along three measured lines for laser sample. (a) . mag — upper line. (b) . mag — middle line. (c) . mag — bottom line. figure 7. comparison of rs obtained by xrd in direction l’ and neutron diffraction in direction t along three measured lines for mag sample. 36 vol. 9/2017 real structure and residual stresses in advanced welds (a) . laser — upper line. (b) . laser — middle line. (c) . laser — bottom line. figure 8. comparison of fwhm obtained by xrd in direction l’ and neutron diffraction in direction t along three measured lines for laser sample. (a) . mag — upper line. (b) . mag — middle line. (c) . mag — bottom line. figure 9. comparison of fwhm obtained by xrd in direction l’ and neutron diffraction in direction t along three measured lines for mag sample. 37 k. trojan, ch. hervoches, k. kolařík et al. acta polytechnica ctu proceedings references [1] t. nitschke-pagel, k. digler. sources and consequences of residual stresses due to welding. mater sci forum pp. 2777–2785, 2014. doi:10.4028/www.scientific.net/msf.783-786.2777. [2] m. sokolov, a. salminen, m. kuznetsov. laser welding and weld hardness analysis of thick section s355 structural steel. mater design 32(10):5127–5131, 2011. doi:10.1016/j.matdes.2011.05.053. [3] c. lachmann, t. h. nitschke-pagel, h. wohlfahrt. nondestructive characterization of residual stress relaxation and fatigue processes in cyclically loaded welded joints. pp. 1400–1409. oxford, uk, 2000. [4] w. woo, et al. neutron diffraction measurements of residual stresses in a 50 mm thick weld. mater sci eng a 528(12):4120–4124, 2011. doi:10.1016/j.msea.2011.02.009. [5] c. randau, u. garbe, h. g. brokmeier. stresstexturecalculator: a software tool to extract texture, strain and microstructure information from area-detector measurements. j appl cryst 44(3):641–646, 2011. doi:10.1107/s0021889811012064. [6] n. jia, et al. an in situ high-energy x-ray diffraction study of micromechanical behavior of multiple phases in advanced high-strength steels. acta mater 57(13):3965– 3977, 2009. doi:10.1016/j.actamat.2009.05.002. 38 http://dx.doi.org/10.4028/www.scientific.net/msf.783-786.2777 http://dx.doi.org/10.1016/j.matdes.2011.05.053 http://dx.doi.org/10.1016/j.msea.2011.02.009 http://dx.doi.org/10.1107/s0021889811012064 http://dx.doi.org/10.1016/j.actamat.2009.05.002 acta polytechnica ctu proceedings 9:32–38, 2017 1 introduction 2 experimental methods 3 experiment 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0107 acta polytechnica ctu proceedings 22:107–112, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app chloride resistance of nano-particle repaired concrete rostislav šulc∗, michal himmel, jiří němeček czech technical university in prague, faculty of civil engineering, department of construction technology, thákurova 7/2077, 166 29 prague 6, czech republic ∗ corresponding author: rostislav.sulc@fsv.cvut.cz abstract. this paper describes the results of the experimentally modeled chloride ion penetration into different quality concrete structures with fully developed carbonation. parts of concrete samples were further treated, to increase the resistance of concrete to chloride ion penetration, with a colloidal solution of nano-silica or solution of sodium methyl siliconate. measurement of chloride profiles was performed on samples subjected to free diffusion of cl− ions and on samples subjected to accelerated tests. the chloride ion profile in the samples was determined in layers of 5 mm thickness up to a total of 45 mm. [1, 2] keywords: chloride resistance, nano-particle, nano-silica, concrete resistance, chloride accelerated test, natural diffusion. 1. introduction the motivation for this experiment is to verify the possibilities of using nano-sized particles to close the porous structure of the concrete. therefore, a colloidal solution nano-silica (ns) or 10% solution of sodium methyl siliconate (sms) were used for post-mix treatment of concrete. the solid particles of the solution fill up and clog the open pores that are entering points for salt molecules into the concrete. such a closed structure should better withstand aggressive salt water and thereby increase the resistance of concrete to counteract the action of chlorides dissolved in water. the result of this experiment is to provide information on the amount of chlorides penetrating into the depth of the concrete and to limit the possibility of reducing the amount of chlorides entering the concrete structure by closing the microstructure of the concrete. for this test were the used mixtures of concrete also fully carbonated. [1–4] 2. materials and mix proportions concrete specimens were made from portland cement (cem i 42.5 r location mokrá) as a binder and sand and natural crushed aggregate from zbraslav location were used as aggregate. no other admixtures or additives were used. in table 1 is shown mixture composition per 1 m3. for this experiment was prepared mixture so called “normal concrete” (no2a) quality concrete with predicated compressive strength over 40 mpa. second was used so called “low quality concrete” with compressive strength around 20 mpa. concrete was mixed for 5 minutes in laboratory concrete mixer, subsequently was mixture cast into cylindrical and cubic moulds and vibrated. the samples were unmoulded and submerged into water bath after 48 hours for mixing. testing cubes of dimensions 150×150 ×150 mm were used for compressive strength test and from cylinders with 100 mm in diameter and height of 200 mm were prepared test samples with a diameter of 100 mm and a height of 50 mm for the chloride migration tests. measured compressive strength of testing concrete samples was 48.9 mpa for concrete no2a. this value was determined as the average of three test specimens. the average bulk density of the concrete was 2351 kg·m−3. this value was also determined as the average of three test specimens. for concrete lo2 these numbers are: 20.5 mpa and 2230 kg·m−3. after cutting of samples was done, the carbonation process began. samples were left in storage box with moisture around 65% and slow income of gas carbon dioxide co2. until the phenolphthalein test indicated that the samples were fully carbonated. 3. used methods 3.1. vacuum penetration all samples were stored in desiccator in vacuum for minimum of 2 hours. after this was done, distilled water was poured in, so the samples were fully submerged, and after than vacuum was done again in desiccator. thus, the samples were penetrated with water in vacuum for another minimum of 2 hours. [4] 3.2. natural chloride penetration the samples were coated on the underside and the sides with the epoxy adhesive. than the samples were placed in the mold base for the test of natural diffusion. subsequently, vacuum penetration as described above, was performed on the samples. next, a collar was enclosed in the mold and the samples were sealed from all sides except the upper surface. then the samples were submerged in 3% nacl solution and, 107 http://dx.doi.org/10.14311/app.2019.22.0107 http://ojs.cvut.cz/ojs/index.php/app r. šulc, m. himmel, j. němeček acta polytechnica ctu proceedings binder aggregate water ratio cem i-42.5r 0-4 mm 4-8 mm 8-16 mm water/binder (kg/m3) no2a 436.4 872.7 581.8 290.9 186.0lo2 261.8 1150 581.8 291 210.2 table 1. definition of concrete composition. silica [wt.%] 50 ph 9.5 viscosity [cp] 8 density [g·cm−3] 1.4 table 2. typical properties of levasil cb8 [5]. cas no. 16589-43-8 einecs no. 240-648-3 formula ch6o3si·xna flash point >100 °c color and appearance colorless clear liquid solid contents 30% silicone solids contents 19% ph value 12-13 density 25/25 °c 1.25 table 3. typical properties of sodium methyl siliconate. top of mold was sealed to eliminate evaporation, and kept in constant laboratory temperature (22 ± 2 °c) for approximately 150 days. after these workflows were done, the samples were dried in heat chamber at 55 °c for at least 7 days. 3.3. protection by dispersion of colloidal silica same as written above (accelerated migration) with exception of using nanoparticle solutions to the side of cathode. these solutions were levasil cb8 (characteristics of levasil cb 8 are shown in table 2) and sodium methyl siliconate (characteristics are shown in table 3). [4] 3.4. nano-silica levasil cb8 colloidal solution was used in this experiment. levasil cb8 is an alkaline, aqueous dispersion of colloidal silica that is approximately 50% solids by weight. the silica dispersion is sodium stabilized and the amorphous silica particles carry a negative surface charge. the sio2 particles are discrete, have a smooth, spherical shape, and are present in a wide particle size distribution. the physical appearance of the dispersion is a white liquid, slightly more viscous than water. [5] 3.5. sodium methyl siliconate this solution is commonly used as hydrophobic protective layer of surfaces of natural stone, like sandstone, limestone etc. in our experiment the solution was diluted to 10% of solid contents. 3.6. natural chloride penetration after protection by dispersion of colloidal silica samples prepared according to the procedure described under protection by dispersion of colloidal silica were used for this experiment. the penetrated samples were then subjected to the procedure described in natural chloride penetration. the side of the sample, which was exposed to colloidal silicon solution, was then exposed to a solution of 3% nacl. of course, samples were kept in constant laboratory temperature (22 ± 2 °c) for approximately 150 days. after that the samples were dried in heat chamber at 55 °c for at least 7 days. 3.7. electro migration (accelerated migration) after protection by dispersion of colloidal silica test started same way as chloride electro migration. the surface of test sample, which was covered with a protective layer, was putted into this chamber with solution of 3% nacl. into the chamber on the other side of the sample was solution of 0.3 m na(oh) added. again, electrolytes were connected (to the anode side with 0.3 molar naoh solution and to the cathode side with 3% nacl solution) to a power source with a constant voltage of 20 v for 48 hours. [4] 3.8. description of chloride profile measurement on the prepared samples, the chloride profile and the amount of chlorides in the given concrete layer were measured. the chloride concentration was analysed in 5 mm depth-steps. it means, that the concrete specimens were drilled in 5 mm steps and the powder from drilling was collected in amount of 1.5 g. the samples contained material from a depth of 0-5 mm; 5-10 mm; 10-15 mm; 15-20 mm, 2025 mm; 25-30 mm; 30-35 mm; 35-40 mm; 40-45 mm. the results show the values of “average” depth, i.e. 2.5 mm; 7.5 mm; 12.5 mm; 17.5 mm; 22.5 mm; 27.5 mm; 32.5 mm; 37.5 mm; 42.5 mm. the scheme of the compared experiments is shown in table 4. [4] approximately 10 ml of 20% hno3 was added to the collected 1.5 g powder samples and the sample was shaken carefully for approx. for 1 hour and then let to stabilize for approx. 24 hours. [4] 108 vol. 22/2019 chloride resistance of nano-particle repaired concrete figure 1. electro migration chamber. samples no2a with ns protection natural penetrationaccelerated migration with sms protection natural penetrationaccelerated migration without protection natural penetration samples lo2 with ns protection natural penetrationaccelerated migration with sms protection natural penetrationaccelerated migration without protection natural penetration table 4. scheme of experiments. measurement of the concentration of chloride ions in the leaching solution was performed using the ion selective electrode and reference mercury sulphate electrodes scanlab system. prior to each measurement, the ion-selective electrode was calibrated in concentrated kcl solution. a total of 6 calibration solutions of 10-4 were prepared with concentration of 10−3; 3 × 10−3; 10−2; 3 × 10−2; 10−1. the calibration curve was set by breaking the values measured in the calibration solutions with a determination coefficient higher than 0.999. [4] 4. results and discussion all resulting values in all figures are the average values of the measurements made on 3 independent samples. figure 2 shows an electrical current waveform pattern in accelerated migration of chloride ion samples over a period of 48 hours. from the current flow in samples protected by dispersion of colloidal silica, it is obvious that the conductivity of the colloidal solution is lower, but also that the colloidal silicon particles in the solution have a truly negative charge. it is clear from this that the amount of chloride ion transferred was less with protection of levasil cb8. as above in figure 2 with the difference of used concrete. figure 3 shows “normal quality concrete”. with this concrete the difference between used protective layer is extremely small. figure 4 shows the chloride profile in normal quality samples after procedure of natural chloride penetration. in comparison to unprotected samples, protection of levasil cb8 shows zero level of protection, even the protected samples show higher concentrations of chloride ions that unprotected ones. the real question is why sms protected samples shows this kind of graph and not the hyperbole as was expected. figure 5 shows the concentration of chloride ions in the samples of low quality concrete. these samples were subjected to protection of colloidal silica and natural chloride penetration. all of those results the course of the chloride profile corresponds to the total periodic hyperbole that was expected for this type of test. unlike the figure 4 results we can say that the solution of sms works to some degree. absolute results show larger numbers that normal quality concrete but that was quite the point of those mixtures. figure 6 shows the chloride profile in normal quality samples with protective layer of colloidal silica. in comparison there are samples after natural diffusion and accelerated migration tests. solution of sodium methyl siliconate shows better results with protection. in figure 7 is shown the concentration of chloride ions in the samples of low quality concrete. samples subjected to chloride electro migration show significantly lower chloride ion values than naturally chloride penetration samples. the course of the chloride electro migration profile in the chloride electro migra109 r. šulc, m. himmel, j. němeček acta polytechnica ctu proceedings figure 2. the course of the electrical current in the electrolyte during the preparation of the samples. figure 3. the course of the electrical current in the electrolyte during the preparation of the samples. figure 4. concentration of chloride ions normal quality concrete, natural diffusion. 110 vol. 22/2019 chloride resistance of nano-particle repaired concrete figure 5. concentration of chloride ions low quality concrete, natural diffusion. figure 6. concentration of chloride ions normal quality concrete, natural diffusion. 111 r. šulc, m. himmel, j. němeček acta polytechnica ctu proceedings figure 7. concentration of chloride ions low quality concrete, natural diffusion. tion sample suggests that only a very small amount of chloride ions were penetrated into the sample compared to natural chloride penetration. 5. summary the results presented in this paper show the possibility of protection of concrete structures against the effect of chlorides dissolved in water. many experiments have been carried out, which have demonstrated the possibility of simulating the acceleration of the chloride ion penetration into the concrete structure. further experiments were performed showing the possibility of penetration of the concrete structure of micro particles. the results here are to be further checked on real-world constructions. the main points can be summarized as follows: (1.) the chloride ion penetration into concrete can be simulated by passing the electric current through the electrolyte. (2.) the simulated phenomenon is highly dependent on boundary conditions, in particular the strength of the electric current, effect of the time of the electric field and the concentration of the electrolyte solution. (3.) the dispersion of colloidal silica can be penetrated by means of an electric current into concrete samples. (4.) dispersion of colloidal silica can have positive results when closing the porous concrete structure. (5.) dispersion of colloidal silica can reduce the amount of chloride ions dissolved in water when they penetrate into the concrete structure. acknowledgements financial support of the czech science foundation (project 17-05360s) and the grant agency of the czech technical university in prague (sgs16/135/ohk1/2t/11) is gratefully acknowledged. references [1] r. šulc, m. himmel, j. němeček. chloride resistance of nanoparticle repaired concrete measured by accelerated tests. in special concrete and composites 2017, vol. 760 of key engineering materials, pp. 91–95. trans tech publications, 2018. doi:10.4028/www.scientific.net/kem.760.91. [2] p. hlaváček, j. němeček. accelerated chloride migration tests in concrete. in engineering mechanics 2016 book of full texts, pp. 1–4. institute of thermomechanics, as cr, v.v.i., prague, 2016. [3] j. němeček, p. hlaváček, m. kouřil, t. jamborová. electrochemical injection of nanoparticles into concrete. in proceedings of the rilem conference. materials, systems and structures in civil engineering 2016, segment on electrochemistry in civil engineering, pp. 81–87. rilem publications s.a.r.l., prague, 2016. [4] j. němeček, y. xi. nanotechnology in construction, chap. electrochemical injection of nanoparticles into existing reinforced concrete structures, pp. 213–218. springer, new york, 2015. [5] n. akzonobel. levasil cb8 colloidalsilica, data sheet. https://colloidalsilica.akzonobel.com/ construction/concrete/. 112 http://dx.doi.org/10.4028/www.scientific.net/kem.760.91 https://colloidalsilica.akzonobel.com/construction/concrete/ https://colloidalsilica.akzonobel.com/construction/concrete/ acta polytechnica ctu proceedings 22:107–112, 2019 1 introduction 2 materials and mix proportions 3 used methods 3.1 vacuum penetration 3.2 natural chloride penetration 3.3 protection by dispersion of colloidal silica 3.4 nano-silica 3.5 sodium methyl siliconate 3.6 natural chloride penetration after protection by dispersion of colloidal silica 3.7 electro migration (accelerated migration) after protection by dispersion of colloidal silica 3.8 description of chloride profile measurement 4 results and discussion 5 summary acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.16.0018 acta polytechnica ctu proceedings 16:18–24, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app the load-displacement behaviour of ground anchors in fine grained soils juraj chalmovský∗, lumír miča brno university of technology, faculty of civil engineering, institute of geotechnics, veveří 331/95, 602 00 brno, czech republic ∗ corresponding author: chalmovsky.j@fce.vutbr.cz abstract. ground anchors represent an important structural element in the area of geotechnical engineering. despite their extensive usage, a design process of these elements is usually performed using simple empirical and semi-empirical methods, neglecting several important influencing factors. this paper gives an analysis of the factor of non-uniform distribution of skin friction resulting in a progressive failure of ground anchors. first, the finite element method in combination with a material model involving regularized strain softening is utilized. next, an experimental program, including several investigation anchor load tests, was carried out. the goal of this program was to confirm preliminary conclusions drawn from numerical studies and to obtain relevant data for further back analysis. after, there is then described a newly developed application based on the load transfer method, in which all the findings from numerical computations and experimental measurements are incorporated. keywords: ground anchor, strain softening, investigation load test, load transfer function. 1. introduction mobilization of shear stresses along a fixed length of ground anchors is significantly non-uniform during anchor loading. in the first loading stages, shear stresses mobilized at the grout soil interface are concentrated at the top of the fixed anchor length. after reaching the peak value, shear stress is consequently reduced towards its critical and residual value. the location of peak shear stress moves along the fixed anchor length as schematically shown in figure 1. the non-constant (peak) shear stress distribution along the fixed anchor length has been experimentally confirmed by several authors: ostermayer [1], scheele [2], barley [3], woods, and barkhordari [4]. the described phenomenon has a severe influence on the ultimate bearing capacity, ground anchor efficiency and load displacement curves. the ultimate load carrying capacity is not directly proportional to the fixed anchor length. because the maximum bond stress is reached only in a small portion of the fixed anchor length, anchors with fixed lengths longer than 8 m are inefficient (barley [5]). in addition, the assumption of a constant bond stress may lead to an overestimation of the load carrying capacity. based on regression analysis of anchor loading tests performed mainly in london clay and glacial clay, barley [3] has proposed an analytical formula for the efficiency factor fef f and for the load carrying capacity of low pressure grouted anchors tult (eq. 2), where lf ixed is the fixed anchor length, d is the borehole diameter and cu is the undrained soil shear strength. the progressive failure mechanism is dominant in the case of stiff, to very stiff over-consolidated clays and dense, to very dense non-cohesive soil exhibiting peak shear strength followed by a strength reduction during strain softening towards a residual stress state. fef f = 1, 6l−0,57f ixed (1) tult = πdlfef f cu (2) 2. finite element modelling in order to analyze the effect of progressive failure by numerical methods it is necessary to adopt a constitutive model involving post-peak strain softening. for this purpose the multi-laminate model for stiff overconsolidated clays (mlsm) proposed in schadlich [7] was used. mlsm requires 11 input parameters: the oedometric reference stiffness eoed,ref , the modulus of elasticity for unloading reloading eur,ref , the poisson ratio for unloading reloading νur, the shear hardening parameter amat, the critical state friction angle ϕcs, the hvorslev surface inclination ϕe, the initial preconsolidation pressure σnc, the reference pressure pref , the stress dependency index m, the softening scaling factor hsof t, and the internal length lcalc. ground conditions of modeled and tested anchors consist of highly over-consolidated stiff brno clay of the neogene age. svoboda, mašín [8] estimated the value over-consolidation ratio ocr=6.5 based on oedometric tests on non-disturbed samples. applying the formula proposed by mayne, kulhawy [9] leads to k0=1.25. input parameters for mlsm are calibrated based on three cu triaxial tests (svoboda et al. [10]) performed at different confining pressures (275 kpa, 500 kpa, 750 kpa). the calibration process is divided 18 http://dx.doi.org/10.14311/app.2018.16.0018 http://ojs.cvut.cz/ojs/index.php/app vol. 16/2018 the load-displacement behaviour of ground anchors figure 1. non-uniform shear stress distribution along the anchor fixed length [6]. figure 2. triaxial stress strain curves. into two parts described in the following sections. mlsm with determined input parameters is then applied to simulation of a low pressure (gravity) grouted anchor. finally, back analysis of the investigation anchor load test is performed. fem packages plaxis 2d brinkgreve et al. [11] and plaxis 3d brinkgreve et al. [12] were used for the analyses presented here. 2.1. strain hardening parameters first, parameters governing strain hardening plasticity were obtained (table 1). this can be done using a "single stress point approach". soiltestlab application was used for this purpose. comparisons between simulations on a single stress point level and measured data (stress strain curves and effective stress paths) are shown in figure 2 and figure 3. 2.2. strain softening parameters determination of parameters governing strain softening on a single stress point level is not possible, as the softening rate is size dependent and non-homogeneity (shear band) originates in a sample. a 3d model of a triaxial test with real sample dimensions is therefore figure 3. effective stress paths. necessary. the strain softening behaviour is governed by two parameters: the softening scaling factor hsof t which controls the rate of softening for a given damage strain increment and the internal length lcalc which determines the domain to be considered during the regularization. in order to transfer a calibrated softening rate into a boundary value analysis, softening scaling (brinkgreve [13], galavi [14]) can be utilized because the softening rate depends only on the ratio hsof t/lcalc. for the purpose of calibration of these parameters, a 2d plane strain numerical model of a biaxial test and a full 3d numerical model of a triaxial test were employed. measured and predicted stress strain curves for the cu test with confining pressure 275 kpa (closest to the stress range in the anchor vicinity) are shown in figure 4. three simulations of a biaxial test with different hsof t/lcalc ratios (8000, 24000, 32000) were performed. a ratio of hsof t/lcalc=32000 is used in the final prediction with a 3d triaxial test model and for further finite element models. localization of shear strains into the shear band at the end of the simulation is clearly visible in figure 5. in the case of 19 juraj chalmovský, lumír miča acta polytechnica ctu proceedings eoed,ref eur,ref νur amat φcs φe σnc pref m kpa kpa ° ° kpa kpa 1200 8000 0.2 15 20 16 1800 100 0.55 table 1. values of input parameters governing strain hardening. figure 4. triaxial stress strain curves – softening. (a) . deformed mesh (b) . shear band figure 5. post peak behaviour. solving a practical boundary value problem, the internal length lcalc is appropriately selected (according to the recommendations) and hsof t is scaled to get the required hsof t/lcalc ratio. 2.3. initial application in the first step after calibration, an fe – mlsm combination was used for simulating a load test of the gravity grouted anchor with a fixed length lf ixed of 4 m. the axisymmetric finite element model using 6noded triangular elements was prepared. mobilization of skin friction along the fixed length on the groutsoil interface for three different load levels (65 kn, 125 kn, 146 kn) is shown in figure 6. after reaching the peak skin friction of 95 kpa in the second load step, shear stresses are decreasing towards the critical figure 6. shear stress distributions for different loading stages. value. with increasing load level, also the fixed length section in the softening regime is growing. 2.4. back analysis of the investigation anchor load test calculations adopting a simple model of a ground anchor confirmed that it is possible to model the progressive failure mechanism by combination of the finite element method and the constitutive model involving regularized strain softening. in the second step, back analysis of the investigation anchor load test was therefore performed. the investigation load test was carried out by misove [15]. the free and fixed length was 4 m and 8 m, respectively. the anchor was post-grouted by tube-a-manchette system (tam). both section were separated by a packer to prevent a grout inflow into the free length. the anchor was excavated after the test in order to determine its diameter. due to post-grouting, the diameter was enlarged from the original 140 mm to 190 mm. the process of grouting was modelled by applying a pre-calculated volumetric strain to the respective elements thus increasing radial stresses acting on the fixed length surface. three basic calculations were performed: 20 vol. 16/2018 the load-displacement behaviour of ground anchors figure 7. predicted and measured load – displacement curves. • standard calculation with mlsm involving the reaching of peak shear strength (hvorslev surface) followed by strain softening. • calculation with modified version of the mlsm with no peak strength, thus only critical shear strength can be reached. • calculation with the standard mohr – coulomb model with peak shear strength parameters cp=90 kpa and ϕp=14°. these values were determined from the shear box test assuming a normal stress acting on a fixed length surface between 200 and 400 kpa. predicted and measured load – displacement curves are shown in figure 7. the alternative using peak strength characteristics significantly overestimates the bearing capacity. on the other hand calculation with modified mlsm (only critical strength can be reached) underestimates the bearing capacity. a sufficient match in terms of displacements and the carrying capacity is only reached when a skin friction decrease from peak to critical is involved. predicted shear stress distributions for three different load levels (627 kn, 739 kn, and 800 kn) are shown in figure 8. skin friction mobilization is significantly non-uniform. progressive shear stresses decrease from the peak value of 150 kpa and gradually propagate to a distance of 3 m from the fixed length head. 3. experimental program analysis performed using the combination of fem and mlsm has revealed that neglecting a progressive skin friction decrease may lead to highly inaccurate results. it was therefore decided to perform a detailed experimental test program involving 6 investigation load tests (up to a failure) of anchors constructed in brno neogene clay. each anchor was monitored by two groups of systems: figure 8. computed shear stress distributions for three load levels. • standard monitoring: measuring of anchor head displacement (lvdt) and a pre-stressing force. • detailed monitoring: axial strain measuring along the tendon and in grout, using vibrating wire tensometers and electric resistance gauges. in sum 5 of the tested anchors were standard temporary anchors, the last one was sbma (single bore multiple anchor). the free length of standard anchors was 5 m, the fixed length was 6 m (2 tests), 8 m (1 test) and 10 m (2 tests). due to the limited extent of the paper only two representative results are presented. detailed test results and their interpretation can be found in chalmovský [16]. the time record of a measured pre-stressing force and the anchor head displacement is shown in figure 9. on reaching the ultimate bearing capacity (point no. 1) there followed a rapid decrease of pre-stressing force which was accompanied by clearly visible upward movement of the grout body. the anchor was fully unloaded after that (point no. 2) followed by re-increasing of prestressing force. it was found that the bearing capacity dropped to 55% of the original peak value (point no. 3). at this load level anchor head displacements stabilized (point no. 4). similar behaviour was observed on all tested anchors. taking into account that the tendon – grout interface was not broken, it might be concluded that the observed behaviour was a consequence of rapid progressive failure of the grout-soil interface. the second example (figure 10) presented in this paper are measured distributions of pre-stressing force along the fixed length of the anchor (lf ixed=10 m). it is obvious from this figure that for three initial 21 juraj chalmovský, lumír miča acta polytechnica ctu proceedings figure 9. the investigation test record. figure 10. measured distributions of pre-stressing force. load levels (236 kn, 476 kn, 715 kn) only the first 5.5 m of the fixed length is utilized. considering that pre-stressing force of 715 kn is much higher than what would be considered as a design bearing capacity (applying standard design methods), the anchor economic efficiency is low. 4. development of application based on load-transfer functions analysis using the finite element method and experimental results confirmed that the progressive decrease of skin friction along anchor fixed lengths significantly influences an overall load – displacement behaviour and an ultimate bearing capacity of anchors. despite its complexity, numerical analysis with a constitutive model involving regularized strain softening, presents a time demanding task. in order to involve the aspect of a non-uniform shear stress distribution and other factors in a design process, an approach incorporating the load transfer method is therefore proposed. a load transfer function (t-z curve) is the dependence of shear stress mobilized on the surface of a fixed length segfigure 11. assembled load transfer function. ment and its vertical displacement. the load transfer method is frequently adopted for the determination of load settlement curves of vertically loaded piles (coyle, reese [17], coyle and sulaiman [18]). however its use for ground anchors is rare. the important feature (from a practical point of view) of the proposed approach is that the load transfer function shape is determined using standard laboratory tests. no prior loading tests are necessary. the newly developed algorithm includes the following main features: • derivation of load transfer functions from laboratory testing following the procedure stated in kraft et al. [19]. • axial stiffness change due to the occurrence of tensile cracks in the grout material ([20]). • radial stress increases due to the post-grouting. therefore, cylindrical cavity expansion theory (randolph et al. [21]) is adopted. • radial stress decreases due to the grout consolidation (bezuijen, talmon [22], talmon [23]). the investigation load test presented in the chapter 2.4 is re-analyzed using the developed algorithm. three other load tests in two different localities were further used for verification purposes (chalmovský [16]). the load – transfer function is shown on figure 11. the predicted and measured load displacement curves are shown in figure 12. slightly higher displacements are predicted, and a reasonable match is reached for the ultimate bearing capacity. predicted distributions of skin friction and axial displacements for five different loading stages are shown in figure 13 and figure 14. for two last loading stages a shear stress decrease in the front section of the fixed length is already initiated. the tested anchor was equipped by electric resistance gauges placed on the tendon in regular intervals of 1 m. determination of shear stress profiles was therefore possible. figure 14 presents the predicted and measured shear stress profile for the last loading stage. progressive failure propagation is obvious. predicted peak shear stress is in good match with observation. the fixed length section in the softening regime is longer than predicted. 22 vol. 16/2018 the load-displacement behaviour of ground anchors figure 12. computed and measured load displacement curves. figure 13. skin friction distributions. this might be due to slightly higher real bearing capacity (886 kn) which allowed further spreading of the progressive failure. 5. conclusions in order to analyse the non-uniform shear stress distribution along the anchor fixed lengths, three different approaches are presented in the paper: numerical, analytical and experimental. in the first approach the finite element method in combination with the constitutive model involving strain softening was used. the back analysis revealed that ignoring shear stress decrease after peak shear strength is reached may lead to significant bearing capacity overestimation. in-situ loading tests confirmed that the peak and residual bearing capacity should be distinguished. in the presented loading test, the residual peak capacity ratio was 0.55. the newly developed application is based on the load transfer method. load transfer functions can be assembled employing standard laboratory tests. detailed information about all three approaches can be found in chalmovský [16]. figure 14. pre-stressing force distributions. figure 15. comparison of shear stress distribution for last loading stage. list of symbols amat shear hardening parameter [–] cp peak cohesion [kpa] cu undrained cohesion [kpa] d fixed length diameter [m] eoed,ref reference primary oedometer stiffness [kpa] eur,ref reference unloading-reloading stiffness [kpa] fef f coefficient of effectivity [–] hsof t hvorslev surface softening parameter [–] k0 coefficient of earth pressure at rest [–] lcalc internal length for regularization [m] lf ixed fixed length [m] m power for stress dependency of stiffness [–] ocr over-consolidation ratio [–] pref reference stress [kpa] tult bearing capacity [kn] νur poisson ratio for unloading reloading [–] σnc,0 initial pre-consolidation pressure [kpa] φcs critical state friction angle φe inclination of hvorslev surface in τ-σ space φp peak friction angle acknowledgements this research was financially supported by the research project of the ministry of industry and trade no. frti4/329. 23 juraj chalmovský, lumír miča acta polytechnica ctu proceedings references [1] h. ostermayer. construction carrying behaviour and creep characteristics of ground anchors. proceedings diaphragm walls and anchorages, inst civ eng, london pp. 141–151, 1975. [2] f. scheele. tragfahigkeit von verpressankern in nichtbindigen boden neue erkentnisse durch dehnungsmessungen im verankerungsbereich. technische universität münchen, 1981. [3] a. d. barley. the single bore multiple anchor system. proc int conf: ground anchorages and anchored structures, london pp. 65–75, 1997. [4] r. i. woods, k. barkhordari. the influence of bond stress distribution on ground anchor design. proc int conf: ground anchorages and anchored structures, london pp. 55–64, 1997. [5] a. d. barley. theory and practice of the single bore multiple anchor system. anker in theorie und praxis, proc int symposium, salzburg pp. 297–301, 1995. [6] m. puller. deep excavations: a practical manual. 2003. [7] b. schadlich. a multilaminate constitutive model for stiff soils, ph.d. thesis. gruppe geotechnik graz, graz university of technology, austria, 2012. [8] t. svoboda, d. mašín. impact of a constitutive model on inverse analysis of an natm tunnel in stiff clays. proc ita-aites world tunnel congress 2008, agra, india 2:627–636, 208. [9] p. w. mayne, f. h. kulhawy. k0 ocr relationships in soil. proc asce j geotech engdiv 108(6):851–872, 1982. [10] t. svoboda, d. mašín, j. boháč. class a predictions of an natm tunnel in stiff clay. computers and geotechnics 37(6):817–825, 2010. [11] r. b. j. brinkgreve, w. m. swolfs, e. engin. plaxis 2d 2012 users manual. plaxis bv, delft, the netherlands 2012. [12] r. b. j. brinkgreve, w. m. swolfs, e. engin. plaxis 3d 2012 users manual. plaxis bv, delft, the netherlands 2012. [13] r. b. j. brinkgreve. geomaterial models and numerical analysis of softening, ph.d. thesis. delft university press, 1994. [14] v. galavi. a multilaminate model for structured clay incorporating inherent anisotropy and strain softening, ph.d. thesis. gruppe geotechnik graz, graz university of technology, austria, 2007. [15] p. mišove. construction of pres-tressed ground anchors and their bearing capacity (in slovak), ph.d. thesis. vuis, bratislava, 1984. [16] j. chalmovský. behaviour analysis of a ground anchor fixed length in fine grained soils (in czech), ph.d. thesis. brno university of technology, department of geotechnics, 2016. [17] h. coyle, l. c. reese. load transfer for axially loaded piles in clay. journal of the soil mechanics and foundations division 92:1–26, 1966. [18] h. coyle, i. h. sulaiman. skin friction for steel piles in sand. journal of the soil mechanics and foundations division 93(6):261–278, 1967. [19] l. m. kraft, t. kagawa, r. p. ray. theoretical t-z curves. journal of the geotechnical engineering division 107(11):1543–1561, 1981. [20] building code requirements for structural concrete (aci 318) and commentary. [21] m. f. randolph, j. s. steenfelt, c. p. wroth. the effect of pile type on design parameters for driven piles. proc seventh european conference on soil mechanics and foundation engineering 2 pp. 107–114, 1979. [22] a. bezuijen, a. talmon. grout pressures around a tunnel lining, influence of grout consolidation and loading on lining. proc underground space for sustainable urban development (ita singapore) 2004. [23] a. talmon, a. bezuijen. simulating the consolidation of tbm grout at noordplaspolder. tunnelling and underground space technology 24(5):493–499, 2009. 24 acta polytechnica ctu proceedings 16:18–24, 2018 1 introduction 2 finite element modelling 2.1 strain hardening parameters 2.2 strain softening parameters 2.3 initial application 2.4 back analysis of the investigation anchor load test 3 experimental program 4 development of application based on load-transfer functions 5 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0007 acta polytechnica ctu proceedings 22:7–11, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app fresh state properties of lime mortars with flax oil admixture pavla bauerováa, b, pavel reitermana, c, milena pavlíkováb, magdalena kracík štorkánovád, martin kepperta, b, ∗ a czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic b czech technical university in prague, faculty of civil engineering, department of materials engineering and chemistry, thákurova 7, 166 29 praha 6, czech republic c czech technical university in prague, faculty of civil engineering, experimental center, thákurova 7, 166 29 praha 6, czech republic d art & craft mozaika, z.s., kapitulní 103/19, 252 62 únětice, czech republic ∗ corresponding author: martin.keppert@fsv.cvut.cz abstract. flax oil has been used as mortar improving, hydrophobic, admixture already in ancient times. the paper describes the identification of flax oil in mosaic mortar from ca. 1900 by help of ftir spectroscopy. this historic mortar was reproduced by nowadays raw materials and the influence of flax oil on consistency and air entraining of fresh mortar was tested. it was found that already small amount of oil caused significant air entrainment, which, simultaneously with water repellency action, may contribute to higher durability of mortar with oil admixture. the flux oil influenced also consistency of mortar; its presence caused higher flow value of mortar. moreover, the introduction of the paper summarizes principal knowledge about action of natural admixtures in lime mortars and plasters on basis of current literature. keywords: modern mosaics, flax oil admixture, consistency of fresh mortar, hydrophobic admixture. 1. introduction mortar, plaster or even concrete on basis of lime have been used for several thousand years and most likely will be used also in future [1]. the lime based binders offer important advantages – high accessibility of raw materials, simple production technology and consequently low production costs. on the other hand, lime has obviously numerous disadvantages. one has to distinguish air lime (containing predominantly cao/mgo) or hydraulic lime. air lime is obviously non-hydraulic binder, which has to be in contact with air to set and harden. hydraulic lime is able to reach higher strength and better durability [2]; it can be prepared by calcination of limestone with proper content of hydraulic oxides (nhl) or by mixing of lime with a pozzolanic additive (hl). all of these materials have been modified by numerous natural admixtures since ancient times in order to improve the binder performance – repellency to water, plasticity, rate of setting etc. [3]. moreover, organic fibrous materials have been used as dispersed reinforcement reducing volume changes and increasing strength and integrity of material [4]. nowadays, the application of synthetic organic admixtures (i.e. better defined, more effective, with higher stability) is dominant in engineering practice, nevertheless the studying of natural admixtures has an importance with respect to restoration works or possibly may contribute to higher sustainability of building structures. the natural species used historically as lime admixtures can be classified, according to their chemical nature, to three groups: proteins, polysaccharides and triglycerides (glycerin esters of fatty acids, i.e. oils and fats,). specifically, protein admixtures were caseins (curd, buttermilk, whey), collagens (animal glue) and egg proteins [5]. animal glue was found to increase the mortar’s strength significantly; the mechanism probably lies in refinement of pore system and reduction of total porosity [3]. egg was used as plasticizer and setting retarder [6]; egg white (optimum content 6% by lime) increased plasticity, strength and reduce water absorption [7]. casein containing admixtures improve workability and strength; casein forms in lime environment calcium caseinate, which is insoluble and acts as binder [8, 9]. animal blood can be also classified as protein admixture, even though its complex composition; it is significantly improving adherence of lime mortar and its durability due to reduction of water absorption and air-entraining effect [10]. the ancient polysaccharide admixture is opuntia powder (or mucilage). it increases the rate of carbonation of lime, what obviously results to faster hardening of mortar [3]. nowadays, derivatives of natural polysaccharides cellulose ethers, welan gum, diutan gum, guar gum, xanthan gum are widely used as viscosity-enhancing admixtures, increasing viscosity, cohesion and stability of mortars [11]. influence of starch on lime mortar was found to be dependent on its 7 http://dx.doi.org/10.14311/app.2019.22.0007 http://ojs.cvut.cz/ojs/index.php/app p. bauerová, p. reiterman, m. pavlíková et al. acta polytechnica ctu proceedings dosage; it acted as a thickener when the incorporated dosage was up to 0.30% of lime weight; conversely, above that dosage, it behaved as a plasticizer [12]. triglycerides admixtures are represented e.g. by olive oil; it reduced the porosity as well as pore size resulting to lower permeability of mortar [3]. thanks to their structure, triglycerides are also acting as hydrophobic admixture. linseed oil acted as absorption reducer (and durability enhancer), but without a large influence on pore system [13]. nowadays, the fatty acids are used as water repellant admixtures in form of soaps (e.g. calcium stearate) [14]. the present paper deals with influence of flax (linseed) oil on fresh state properties of lime based mortar. the direct motivation for this research lies in fact, that flax oil was identified in a sample of historical mosaic mortar; the flax oil identification is described as well. 2. experimental the initial material under study, inspiring the present experimental work, was a sample of mosaic mortar collected at peluněk sepulcher, malvazinky cemetery, prague. it was built round 1900. the sample of mortar was characterized by xrd (panalytical empyrean, cukα source); the content of filler insoluble in acid (i.e. quartz and silicates) was determined by help of 1:3 hcl dissolution and filtration. the historic mortar was also examined by sem (zeiss merlin). the oil admixture was identified by help of ftir spectroscopy (nicolet 6700) in extract of mortar (extraction 1:10 by chloroform). the reproduced mosaic mortar was prepared on basis of analysis of the historical mortar from sepulcher. lime hydrate was used as the principal binder (40%), as the filler were used three fractions: quartz 0-0.5 mm (30%) and 0.5-1 mm (10%) and milled limestone d8 (20%). the water/lime ratio was kept constant (0.87). the natural flax oil was dosed as admixture (0-10% by lime hydrate weight). the first step of mixing procedure [15] was homogenization of dry powders by help of electric mixer; than small amount of powder was mixed separately with oil and this paste was dispersed in the rest of dry powders. finally the water was added. the consistency and entrained air content of mortar was measured by standard procedures (en 1015-3 and 7) by help of flow table test and pressure type air content meter. 3. results and discussion characterization of historic mortar: the sample of historic mosaic mortar was examined by xrd (fig. 1). it revealed that the main phases are quartz and calcite; the used filler also introduced muscovite and albite to the mortar. surprisingly, in sample of age of 120 years, also portlandite was found. its presence was confirmed by thermal analysis as well. the results of thermal analysis also indicated, that the binder was closer to the air lime than to the hydraulic one, although edx analysis indicated some si and al atoms dispersed in binder phase. the residuum after acid dissolution of the sample was 15%; this amount thus may be attributed to siliceous filler. unfortunately, the amount of historic mortar sample was limited (ca 10 g) due to conservation of historic mosaic and thus this result cannot be assumed as fully representative; neither the granulometry of filler cannot be determined reasonably. nevertheless, visually one can sate that the mortar does not contain particles larger than 1 mm (these would, obviously, make difficult the application of mosaic mortar). the sem (and edx) analysis shown, that besides siliceous filler, the mortar contains also calcite (marble) flour with size of particles in the range of a few of tenths microns (fig. 2). the carbonate dissolution liberated not only the siliceous aggregates, but also an unknown organic substance forming a brownish greasy layer on the filter paper. in order to identify this substance, 5 g of mortar powder was extracted by chloroform (50 ml). the obtained extract was analyzed by ftir spectroscopy (fig. 3). the spectrum was compared with nowadays flax stand oil. the group of three signals around 2900 cm−1 belongs to c=c groups; in this case in fatty acids. flax oil triglycerides contain three unsaturated fatty acids triply unsaturated α-linolenic acid, doubly unsaturated linoleic acid and monounsaturated oleic acid. the peak at 1758 cm−1 in reference stand oil corresponds to c=o in esters (i.e. triglycerides); the shift 1709 cm−1 in extract is due to decomposition of triglycerides to ca-soaps of carboxylic acids and glycerol in basic environment of mortar. the peak at 1465 cm−1 belongs to –ch2– groups in acids, while the peaks in range 1000-1400 cm−1 are due to c-o bonds in glycerol. the ftir results can be understood as proof that flux oil was used as admixture in historic mosaic mortar. the quantification of organic substance in mortar was attempted on base of mass balance of the extraction experiment; the result was ca. 6% what seems to be a lot and reasonability of such high dosage of flax oil must be checked in future experiments with reproduced mortars. influence of flax oil on fresh state properties of reproduced mortar was tested in wide range of oil dosing (0-10% by lime weight) (fig. 4). the consistency of reference mortar was 130 mm, what is less than 140-200 mm, what is range of “plastic mortar”. in fact, the mortar is usually mixed for 160 mm consistency. nevertheless, this initial (low) choice of mixing water amount was due to possibility to study the oil influence on consistency. the oil dosage generally increased the flow value, but even the very high dosage (10%) did not caused any unacceptable behavior of fresh mortar. the observed trend is opposite compared to publications [8, 15] where the reduced flow is explained by hydrophobicity of the oil and consequent low absorption of water to particles. in our opinion, this effect should on contrary increase the flow value 8 vol. 22/2019 fresh state properties of lime mortars with flax oil admixture figure 1. x-ray diffractogram of historic mortar. figure 2. sem of historic mosaic mortar. (at constant water content), since the mixing water is less adsorbed by particles of lime and thus reduces more effectively the mortar stiffness. the air content of reference mortar was ca. 2%, what is expectable value for lime mortar. already 0.5% of flax oil doubled the air content, while the further increasing doses did not have such significant effect and the value remained around 4-5%. as the air entraining admixtures are conventionally used various surfactants [16], nevertheless also oils are able to increase the air content and slump of concrete [17]. the content of entrained air in a binder is closely related to the viscosity – higher air content means lower viscosity [18], i.e. higher result of flow table test, as it was indicated by the present results. 4. conclusions flax oil admixture was identified by ftir spectroscopy in the sample of historic mortar from peluněk sepulcher, built around 1900 in prague malvazinky cemetery. the composition of this historic mortar was reproduced; the effect of wide range of dosing of flax oil on fresh state properties of mortar was tested experimentally. the oil admixture caused increase of flow value as well as entrained air content. these findings are in agreement with results published on behavior of oil in concrete. moreover the oil is inducing hydrophobic behavior of the mortar. the higher air content and hydrophobicity of mortar are supposed to be responsible for the improved durability of flax oil mortars reported in literature. 9 p. bauerová, p. reiterman, m. pavlíková et al. acta polytechnica ctu proceedings figure 3. ftir spectra of extract, reference oil and extraction medium. figure 4. influence of flax oil admixture on consistency (flow table test) and air content of reproduced mortars. acknowledgements the work has been supported by czech science foundation project nr. 18-13525s “modern mosaic mortars in a microscope – methods for their materials characterization and degradation studies”. references [1] m. stefanidou, v. pachta, s. konopissi, et al. analysis and characterization of hydraulic mortars from ancient cisterns and baths in greece. materials and structures 47:571–580, 2014. doi:10.1617/s11527-013-0080-y. [2] k. dvořák, i. hájková. the effect of high-speed grinding technology on the properties of fly ash. materiali in tehnologije 50:683–687, 2016. doi:10.17222/mit.2015.127. [3] l. ventolà, m. vendrell, p. giraldez, l. merino. traditional organic additives improve lime mortars: new old materials for restoration and building natural stone fabrics. construction and building materials 25:3313– 3318, 2011. doi:10.1016/j.conbuildmat.2011.03.020. [4] c. kennedy, w. a revie, l. troalen, et al. studies of hair for use in lime plaster: implications for conservation and new work. polymer degradation and stability 98:894– 898, 2013. doi:10.1016/j.polymdegradstab.2013.01.004. [5] i. krizova, j. schultz, i. nemec, et al. comparison of analytical tools appropriate for identification of proteinaceous additives in historical mortars. analytical and bioanalytical chemistry 410:189–200, 2018. doi:10.1007/s00216-017-0709-8. [6] r. nogueira, a. p. ferreira pinto, a. gomes. design and behavior of traditional lime-based plasters and renders. review and critical appraisal of strengths and weaknesses. cement and concrete composites 89:192– 204, 2018. doi:10.1016/j.cemconcomp.2018.03.005. [7] m. a. othuman mydin. preliminary studies on the development of lime-based mortar with added egg white. 10 http://dx.doi.org/10.1617/s11527-013-0080-y http://dx.doi.org/10.17222/mit.2015.127 http://dx.doi.org/10.1016/j.conbuildmat.2011.03.020 http://dx.doi.org/10.1016/j.polymdegradstab.2013.01.004 http://dx.doi.org/10.1007/s00216-017-0709-8 http://dx.doi.org/10.1016/j.cemconcomp.2018.03.005 vol. 22/2019 fresh state properties of lime mortars with flax oil admixture international journal of technology 8(5):800–810, 2017. doi:10.14716/ijtech.v8i5.442. [8] p. rovnaníková. omítky. stop praha, 2002. [9] s. kuckova, m. krizkova, l. vankova, et al. towards proteomic analysis of milk proteins in historical building materials. international journal of mass spectrometry 284:42–46, 2009. doi:10.1016/j.ijms.2009.01.011. [10] s. q. fang, k. zhang, h. zhang, b. zhang. a study of traditional blood lime mortar for restoration of ancient buildings. cement and concrete research 76:232–241, 2015. doi:10.1016/j.cemconres.2015.06.006. [11] m. vyšvařil, t. žižlavský. effect of chitosan ethers on fresh state properties of lime mortars. iop conference series: materials science and engineering 251:012039, 2017. doi:10.1088/1757-899x/251/1/012039. [12] a. izaguirre, j. lanas, j. álvarez. behaviour of a starch as a viscosity modifier for aerial lime-based mortars. carbohydrate polymers 80:222–228, 2010. doi:10.1016/j.carbpol.2009.11.010. [13] c. nunes, z. slížková. freezing and thawing resistance of aerial lime mortar with metakaolin and a traditional water-repellent admixture. construction and building materials 114:896–905, 2016. doi:10.1016/j.conbuildmat.2016.04.029. [14] a. izaguirre, j. lanas, j. álvarez. effect of waterrepellent admixtures on the behavior of aerial lime-based mortars. cement and concrete research 39:1095–1104, 2009. doi:10.1016/j.cemconres.2009.07.026. [15] c. nunes, z. slížková. hydrophobic lime based mortars with linseed oil: characterization and durability assessment. cement and concrete research 61-62:28– 39, 2014. doi:10.1016/j.cemconres.2014.03.011. [16] g. liu, l. chen. development of a new type of green switch air entraining agent for wet-mix shotcrete and its engineering application. advances in materials science and engineering 2016:1–9, 2016. doi:10.1155/2016/6981091. [17] b. hamad, a. a. rteil, m. el-fadel. effect of used engine oil on properties of fresh and hardened concrete. construction and building materials 17:311–318, 2003. doi:10.1016/s0950-0618(03)00002-3. [18] f. cardoso, v. john, r. pileggi. rheological behavior of mortars under different squeezing rates. cement and concrete research 39:748–753, 2009. doi:10.1016/j.cemconres.2009.05.014. 11 http://dx.doi.org/10.14716/ijtech.v8i5.442 http://dx.doi.org/10.1016/j.ijms.2009.01.011 http://dx.doi.org/10.1016/j.cemconres.2015.06.006 http://dx.doi.org/10.1088/1757-899x/251/1/012039 http://dx.doi.org/10.1016/j.carbpol.2009.11.010 http://dx.doi.org/10.1016/j.conbuildmat.2016.04.029 http://dx.doi.org/10.1016/j.cemconres.2009.07.026 http://dx.doi.org/10.1016/j.cemconres.2014.03.011 http://dx.doi.org/10.1155/2016/6981091 http://dx.doi.org/10.1016/s0950-0618(03)00002-3 http://dx.doi.org/10.1016/j.cemconres.2009.05.014 acta polytechnica ctu proceedings 22:7–11, 2018 1 introduction 2 experimental 3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0083 acta polytechnica ctu proceedings 22:83–87, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app the use of lightweight aggregate in preparation of thermal insulation lime-based renders jaroslav pokorný∗, milena pavlíková, zbyšek pavlík czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jaroslav.pokorny@fsv.cvut.cz abstract. lime-based renders are common part of historical or older buildings that don’t always provide a comfortable inside climate due to the problems with high thermal losses. thermal insulation renders may possibly retrofit or replace original render layers and mitigate negative impacts of harmful external climate. in this respect, determination of basic structural, mechanical and thermal properties of lime-based renders containing various amount of perlite that was used as partial silica sand replacement is presented in the paper. experimental tests performed for 28 days high relative humidity-cured samples showed significant decrease in bulk density and apparent density for renders with incorporated perlite compared to reference render mix composed of silica sand-based aggregate only. accordingly, porosity of perlite mortars was significantly higher, what led to the lower thermal conductivity values in comparison with reference render. although the mechanical resistance of perlite-modified renders was lower than that of reference material, it was still sufficient for their use as thermal insulation layer compatible with older construction materials. based on the obtained data it can be concluded, the analysed hydrated lime-based plasters with perlite admixture can be considered as promising materials for buildings refurbishment. keywords: hydrated lime, perlite, renders, thermophysical properties. 1. introduction from historical point of view, rendering mortars have always been considered and applied as walls and building facades layers mainly serving to the masonry protection against weathering causing its gradual decay. nevertheless, their decorative aspects were also of considerable importance to achieve aesthetical perfection [1, 2]. aerial lime plasters and mortars belong among the traditional building materials used for masonry and facades until late of the 19th century. lime was very often applied binder due to its ability to add plasticity and good workability to mortar mixture [3]. however, setting and hardening of pure lime mortars is a slow process because of its dependence mainly on reaction of calcium hydroxide with atmospheric carbon dioxide as shown in eq.(1) ca(oh)2 + co2 −→ caco3 + h2o. (1) on the other hand, the specific porous structure of lime mortars with high open porosity, usually ranging from 22 to 35%, allows to pass a certain amount of moisture present in the masonry and thus keep the structure breathable [1, 4]. despite the protection and decorative aspects of traditional renders, their thermal conductivity is high compared to the en 998-1 definition of a thermal insulation render (λ ≤ 0.20 w/mk) and varies typically in dry condition from 0.12 to 0.80 w/mk [2, 5]. the energy retrofitting of existing buildings should be considered as a reasonable approach to reduce the overall building energy consumption [6]. as presented by barbero-barrera, facades are, among the thermal envelope components, the constructive system with the highest impact energy demand of buildings [7]. most of buildings are made up of masonry with traditional renders which can be replaced or retrofit by thermal insulation renders made in the same way as those traditional. in production of thermal insulation renders, highly preferred possibility is adding of various types of mineral lightweight aggregates, such as vermiculite, zeolite, diatomite, perlite and pumice [6]. in this connection, abidi et al. [8] tested plasters on gypsum basis containing perlite and vermiculite. experimental results showed decrease of thermal conductivity up to 0.16 w/mk from initial value 0.50 w/mk. corbanaro et al. [9] presented similar observation for mortars containing natural hydraulic lime, portland cement and perlite. with regarding to the number of performed studies aimed at currently used renders, the influence of silica sand substitution with perlite on properties of traditional lime-based render is examined in the paper. performed experiments pointed to the significant lightening of modified renders accompanied with improvement of their thermal insulation function. 83 http://dx.doi.org/10.14311/app.2019.22.0083 http://ojs.cvut.cz/ojs/index.php/app j. pokorný, m. pavlíková, z. pavlík acta polytechnica ctu proceedings material specific surface bulk density apparent density d10 d50 d90 [m2/kg] [kg/m3] [kg/m3] µm hydrated lime 2 211 ± 25 233 ± 12 2 210 ± 11 0.8 4.2 50.3 table 1. characteristics of lime hydrate cl 90-s. substance content in kg/m 3 m-r mp 25 mp 50 mp 75 mp 100 hydrated lime 326.1 336.5 364.6 384.6 408.1 sand 0.0/0.5 434.5 336.5 243.1 128.3 sand 0.5/1.0 434.5 336.5 243.1 128.3 sand 1.0/2.0 434.5 336.5 243.1 128.3 perlite ex 100 2.5 5.3 8.5 11.1 perlite pb 150 41.7 89.2 141.2 187.7 water 391.3 319.7 266.0 200.0 134.3 table 2. proportion of mortar mixes. 2. experimental 2.1. studied materials and samples preparation as a fundamental binder component of mortar mixes, hydrated lime cl 90-s produced in factory vápenka čertovy chody inc. (lhoist s.a. group), czech republic, was used. hydrated lime disposes very high purity and meets requirements of the standard en 459-1 [10]. basic properties and particle size distribution data of cl 90-s are summarized in table 1. a silica sand (filtrační písky ltd., czech republic) composed of three fractions 0.0/0.5; 0.5/1.0 and 1.0/2.0 mm combined together in mass ratio 1/1/1, was applied as reference aggregate. in modified render mixes, sand was partially or fully substituted with two commercially available types of perlite, ex 100 and pb 150, mixed together in volume ratio 1/4. five types of mortar batches were prepared. silica aggregate in reference mix m-r was replaced by lightweight aggregate in the amount of 25; 50; 75 and 100% by volume. the dosage of batch water corresponded to the value of spreading 160 × 160 ± 5 mm that was kept constant for all renders. before adding, perlite mix was left in plastic box filled by water for 24 hours. composition of particular mortar mixes is expressed in table 2. from fresh render mixes, prisms having dimensions 40 × 40 × 160 mm and cubes with side of 70 mm were casted. in samples casting, the filled moulds were compacted by 5 hits on laboratory table. after 48 hours, the particular samples were unmoulded and stored for 26 days in conditions with high relative humidity of approx. 98% and at temperature of 20 ± 1 °c. 2.2. particle size distribution particle size distribution of hydrated lime was accessed using a laser diffraction particle size analyser analysette 22 nanotec (fritsch) equipped with a red and green laser allowing to detect particles in diameter range from 0.08 to 2 000µm. 2.3. sieving analysis determination of particle size distribution of aggregates was performed according to the standard en 933-1 [11]. at first, aggregate sample was dried at 105 ± 5 °c to constant mass. then, after natural cooling, the sample was placed into series of sieves with mesh size 2.0; 1.0; 0.5; 0.25; 0.125 and 0.063 mm respectively, and mechanically shaken by vibratory sieve shaker retsch as 200. 2.4. pozzolanic activity and specific surface area pozzolanic activity was studied for a finer type of perlite only. for its asseesment, modified chapelle test according to the french norm nf p 18-513 was performed [12, 13]. the method is based on the reaction of 1 g of tested powder material with 2 g cao in water [13]. the value of aggregate specific surface was determined according to the standard en 196-6 with the use of blaine apparatus [13, 14]. 2.5. basic physical properties studied renders were characterized by their basic structural properties such as bulk density, apparent density and total open porosity. the bulk density was measured according to the standard en 1015-10 [15] by weighing sample dry mass and determination of its volume by a digital calliper [16]. the apparent density was accessed using automatic helium pycnometer pycnomatic atc (porotec). the total open porosity was calculated on the basis of knowledge of the bulk and apparent density values. the relative expanded uncertainty of applied testing method was 5% [16]. 84 vol. 22/2019 the use of lightweight aggregate. . . figure 1. sieve analysis of perlite and sand mix. 2.6. mechanical properties among the mechanical properties, compressive strength, flexural strength and dynamic young’s modulus were examined. the strength tests were performed according to the standard en 1015-11 [17]. the flexural strength was tested in a three point bending arrangement on standard prisms having dimensions of 40 × 40 × 160 mm. the broken halves from the flexural strength tests were then used for determination of the compressive strength, whereas the loading area was 40 × 40 mm [18]. the relative expanded uncertainty of the both strength tests was 1.4%. the dynamic young’s modulus was measured by the pulse ultrasonic method (astm c597 – 16), using a dio 562 device (starman’s electronics) [19]. the expanded uncertainty of the dynamic young’s modulus determination was 2%. 2.7. thermal properties the measurement of thermal parameters was conducted on a transient impulse method principle using device isomet 2114 (applied precision) equipped with surface circular probe [16]. the dried plaster samples in the form of cubes with side dimension of 70 mm were used for the measurement [16]. additionally, thermophysical parameters of perlite mix, and silica aggregate in dependence on compaction time were tested. for the thermal conductivity (λ) in the range 0.05–2.0 w/mk was the measuring accuracy 10% of reading. for the volumetric heat capacity (c) in the range 4.0×104−4.0×106 j/m3k, the measuring accuracy was 15% of reading [16]. 3. results and discussion material characteristics of applied aggregates are given in table 3. the reason for combination of two types of perlite was on the one hand to add more fine powder and gain smooth particle size distribution curve, as indicates figure 1. on the other hand, perlite ex 100 (finer than pb 100) showed pozzolanic activity that could potentially support creation of a denser composite structure due to formation of hydrates of c-s-h type [20]. thermophysical parameters of perlite mix, as well as silica aggregate in dependence on compaction time, are given in table 4. from obtained data is clearly visible the high thermal conductivity of silica aggregate that is approx. 7.5 times higher compared to perlite. accordingly, the volumetric heat capacity of silica aggregates was approx. 9 times higher than data obtained for perlite mix. with prolongation of compaction time, both the thermal conductivity and volumetric capacity increased as a consequence of minimization of air gaps between neighbouring aggregate grains. the procedure of this test simulated aggregate behaviour during mix casting and compaction. there is evident, after 20 s (sand) or 30 s (perlite mix) the values of the tested thermophysical parameters were practically stable. the basic structural properties of studied plasters are shown in table 5. perlite incorporation resulted in a significant decrease in renders bulk and apparent densities. accordingly, the increase in the total open porosity was quite apparent. these findings we assign to the low apparent density of perlite and porosity of embedded perlite particles. the total substitution of sand with perlite mix resulted in the one-third unit weight of render mp 100 compared to that measured for reference render. in general, data on structural properties were promising in the sense of improvement thermal insulation performance of perlite-modified renders because high porosity significantly retards heat transport due to the low thermal conductivity of dry air. the mechanical resistance data measured on 28 days cured samples is summarized in table 6. as expected, with increasing dosage of lightweight aggregate in render composition and thus with increase in renders porosity, gradual decline in all tested mechanical parameters was observed. the above mentioned assumption of contribution of pozzolanic activity of perlite ex 100 to creation of more-dense render structure was not confirmed, probably due to the low content of this pozzolana active material in the mix and limitation of contact of its particles with lime due to the presence of other fine and coarser aggregate particles. quantitatively, the highest drop in both compressive and flexural strengths under 0.5 mpa exhibited render mp 100, what is in agreement with its highest porosity. thermophysical properties measured on dried hardened cubic samples are given in table 7. silica sand undeniably caused high thermal conductivity of reference render exceeding 1.3 w/mk. on the other hand, influence of porous lightweight aggregate on improvement of thermal insulation function of analysed materials was quite apparent. the higher the perlite content in the render mix, the lower the thermal conductivity of the sample. in fact, the total substitution of silica sand by perlite in tested render mixes resulted in the thermal conductivity even eleven times lower 85 j. pokorný, m. pavlíková, z. pavlík acta polytechnica ctu proceedings material bulk density apparent density pozzolanic activity[kg/m3] [kg/m3] [mg ca(oh)2/1 g] perlite ex 100 (0.0/0.5) 55 ± 5 651 ± 5 959 perlite pb 150 (0.0/2.0) 233 ± 12 572 ± 4 sand 0.0/0.5 1 478 ± 40 2 654 ± 13 sand 0.5/1.0 1 505 ± 45 2 648 ± 13 sand 1.0/2.0 1 568 ± 50 2 640 ±12 table 3. material parameters of used aggregates. time [s] sand mix | perlite mix bulk density λ c× 106 bulk density λ c× 106 [kg/m3] [w/mk] [j/m3k] [kg/m3] [w/mk] [j/m3k] 0 1 657 ± 25 0.41 ± 0.4 1.57 ± 0.24 125 ± 2 0.05 ± 0.01 0.17 ± 0.02 10 1 910 ± 28 0.56 ± 0.6 1.68 ± 0.25 147 ± 3 0.06 ± 0.01 0.19 ± 0.03 20 1 916 ± 29 0.58 ± 0.6 1.68 ± 0.25 151 ± 3 0.06 ± 0.01 0.20 ± 0.03 30 1 922 ± 22 0.58 ± 0.6 1.68 ± 0.25 153 ± 3 0.06 ± 0.01 0.22 ± 0.03 60 1 927 ± 22 0.58 ± 0.6 1.69 ± 0.26 153 ± 3 0.06 ± 0.01 0.22 ± 0.03 table 4. thermophysical properties of sand and perlite aggregate in dependence on compaction time. material bulk density apparent density total open porosity[kg/m3] [kg/m3] [%] m-r 1 756 ± 26 2 598 ± 13 32.4 ± 1.6 mp 25 1 514 ± 21 2 389 ± 11 36.6 ± 1.8 mp 50 1 216 ± 18 2 101 ± 11 42.1 ± 2.0 mp 75 945 ± 13 1 835 ± 9 48.5 ± 2.3 mp 100 617 ± 9 1 378 ± 7 55.6 ± 2.8 table 5. basic structural properties of studied renders. material flexural strength compressive strength young’s modulus[mpa] [mpa] [gpa] m-r 1.1 ± 0.1 1.3 ± 0.1 4.4 ± 0.1 mp 25 0.8 ± 0.1 1.0 ± 0.1 3.3 ± 0.1 mp 50 0.6 ± 0.1 0.9 ± 0.1 1.8 ± 0.1 mp 75 0.5 ± 0.1 0.8 ± 0.1 1.4 ± 0.1 mp 100 0.4 ± 0.1 0.4 ± 0.1 0.7 ± 0.1 table 6. mechanical parameters of studied renders. material thermal conductivity thermal diffusivity volumetric heat capacity[w/mk] ×10−6 [m2/s] ×106 [j/m3k] m-r 1.34 ± 0.13 0.87 ± 0.08 1.58 ± 0.24 mp 25 0.86 ± 0.08 0.55 ± 0.06 1.53 ± 0.22 mp 50 0.48 ± 0.04 0.32 ± 0.02 1.49 ± 0.22 mp 75 0.26 ± 0.03 0.21 ± 0.02 1.42 ± 0.21 mp 100 0.12 ± 0.01 0.19 ± 0.01 0.60 ± 0.09 table 7. thermophysical parameters of studied renders. 86 vol. 22/2019 the use of lightweight aggregate. . . than that of the reference render. on the contrary, the low bulk density of perlite-modified renders reduced their heat storage capacity, as show gradually decreasing volumetric heat capacity values. 4. summary the effect of silica sand substitution by lightweight aggregate on properties of lime-based renders was studied in the paper. due to the low weight of incorporated perlite mix, the studied renders were markedly lightened. despite pozzolanic activity of perlite ex 100, there was recorded decrease in renders mechanical resistance with increasing perlite dosage in render mix. on the other hand, the high porosity of renders resulted in the reduction of heat transport parameters when the measured thermal conductivity of perlitemodified specimens was even eleven times lower than that of the traditional lime-based material. based on the performed tests and obtained data, it can be concluded, perlite represents lightweight thermoinsulating material suitable for renders prepared by traditional approaches and utilized for renewal or retrofitting purposes. acknowledgements the authors gratefully acknowledge the financial support received from the czech science foundation under project no 18-07332s, and by the ctu in prague under project no sgs17/166/ohk1/3t/11. references [1] a. arizzi, h. viles, g. cultrone. experimental testing of the durability of lime-based mortars used for rendering historic buildings. construction and building materials 28(1):807–818, 2012. doi:10.1016/j.conbuildmat.2011.10.059. [2] y. govaerts, r. hayen, m. de bouw, et al. performance of a lime-based insulating render for heritage buildings. construction and building materials 159:376–389, 2018. doi:10.1016/j.conbuildmat.2017.10.115. [3] b. silva, a. f. pinto, a. gomes. influence of natural hydraulic lime content on the properties of aerial lime-based mortars. construction and building materials 72:208–218, 2014. doi:10.1016/j.conbuildmat.2014.09.010. [4] r. nogueira, a. p. f. pinto, a. gomes. design and behavior of traditional lime-based plasters and renders. review and critical appraisal of strengths and weaknesses. cement and concrete composites 89:192– 204, 2018. doi:10.1016/j.cemconcomp.2018.03.005. [5] en 998-1 specification for mortar for masonry – part 1: rendering and plastering mortar. standard, cen, 2016. [6] e. fenoglio, s. fantucci, v. serra, et al. hygrothermal and environmental performance of a perlite-based insulating plaster for the energy retrofit of buildings. energy and buildings 179:26–38, 2018. doi:10.1016/j.enbuild.2018.08.017. [7] m. barbero-barrera, a. garcía-santos, f. neila-gonzález. thermal conductivity of lime mortars and calcined diatoms. parameters influencing their performance and comparison with the traditional lime and mortars containing crushed marble used as renders. energy and buildings 76:422–428, 2014. doi:10.1016/j.enbuild.2014.02.065. [8] s. abidi, b. nait-ali, y. joliff, c. favotto. impact of perlite, vermiculite and cement on the thermal conductivity of a plaster composite material: experimental and numerical approaches. composites part b: engineering 68:392–400, 2015. doi:10.1016/j.compositesb.2014.07.030. [9] c. carbonaro, s. tedesco, f. thiebat, et al. development of vegetal based thermal plasters with low environmental impact: optimization process through an integrated approach. energy procedia 78:967–972, 2015. doi:10.1016/j.egypro.2015.11.034. [10] en 459-1 building lime – part 1: definitions, specifications and conformity criteria. standard, cen, 2002. [11] en 933-1 test for geometrical properties of aggregated – part 1: determination of particle size distribution – sieving method. standard, cen, 2012. [12] nf p 18-513 pozzolanic addition for concrete metakaolin. definitions, specifications and conformity criteria. standard, afnor, 2010. [13] j. pokornỳ, m. záleská, m. pavlíková, z. pavlík. properties of lightweight composite modified by active siliceous admixture. in aip conference proceedings, vol. 1978. aip publishing, 2018. doi:10.1063/1.5043878. [14] en 196-6 methods of testing cement – part 6: determination of fineness. standard, cen, 2010. [15] en 1015-10 methods of test for mortar for masonry part 10: determination of dry bulk density of hardened mortar. standard, cen, 1999. [16] m. pavlíková, l. zemanová, z. pavlík. thermophysical properties of hydrophobised lime plasters the influence of ageing. in aip conference proceedings, vol. 1866. aip publishing, 2017. doi:10.1063/1.4994510. [17] en 1015-11 methods of test for mortar for masonry part 11: determination of flexural and compressive strength of hardened mortar. standard, cen, 2000. [18] m. pavlíková, l. zemanová, j. pokorný, et al. valorization of wood chips ash as an eco-friendly mineral admixture in mortar mix design. waste management 80:89–100, 2018. doi:10.1016/j.wasman.2018.09.004. [19] j. fořt, a. trník, d. čítek, z. pavlík. residual mechanical properties of hybrid fiber reinforced hpc exposed to high temperatures. in key engineering materials, vol. 722, pp. 52–58. trans tech publications, 2017. doi:10.4028/www.scientific.net/kem.722.52. [20] m. pavlíková, l. zemanová, m. záleská, et al. ternary blended binder for production of a novel type of lightweight repair mortarn. materials 12(6), 2019. doi:10.3390/ma12060996. 87 http://dx.doi.org/10.1016/j.conbuildmat.2011.10.059 http://dx.doi.org/10.1016/j.conbuildmat.2017.10.115 http://dx.doi.org/10.1016/j.conbuildmat.2014.09.010 http://dx.doi.org/10.1016/j.cemconcomp.2018.03.005 http://dx.doi.org/10.1016/j.enbuild.2018.08.017 http://dx.doi.org/10.1016/j.enbuild.2014.02.065 http://dx.doi.org/10.1016/j.compositesb.2014.07.030 http://dx.doi.org/10.1016/j.egypro.2015.11.034 http://dx.doi.org/10.1063/1.5043878 http://dx.doi.org/10.1063/1.4994510 http://dx.doi.org/10.1016/j.wasman.2018.09.004 http://dx.doi.org/10.4028/www.scientific.net/kem.722.52 http://dx.doi.org/10.3390/ma12060996 acta polytechnica ctu proceedings 22:83–87, 2019 1 introduction 2 experimental 2.1 studied materials and samples preparation 2.2 particle size distribution 2.3 sieving analysis 2.4 pozzolanic activity and specific surface area 2.5 basic physical properties 2.6 mechanical properties 2.7 thermal properties 3 results and discussion 4 summary acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.14.0034 acta polytechnica ctu proceedings 14:34–41, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app coupled simulation of gas cooled fast reactor fuel assembly with nestle code system filip osusky∗, stefan cerba, jakub luley, branislav vrban, jan hascik institute of nuclear and physical engineering, faculty of electrical engineering and information technology, slovak university of technology in bratislava, ilkovicova 3, bratislava, slovakia ∗ corresponding author: filip.osusky@stuba.sk abstract. the paper is focused on coupled calculation of the gas cooled fast reactor. the proper modelling of coupled neutronics and thermal-hydraulics is the corner stone for future safety assessment of the control and emergency systems. nowadays, the system and channel thermal-hydraulic codes are accepted by the national regulatory authorities in european union for license purposes, therefore the code nestle was used for the simulation. the nestle code is a coupled multigroup neutron diffusion code with thermal-hydraulic sub-channel code. in the paper, the validation of nestle code 5.2.1 installation is presented. the processing of fuel assembly homogeneous parametric cross-section library for nestle code simulation is made by the sequence triton of scale code package system. the simulated case in the nestle code is one fuel assembly of gfr 2400 concept with reflective boundary condition in radial direction and zero flux boundary condition in axial direction. the results of coupled calculation are presented and are consistent with the gfr 2400 study of the gofastr project. keywords: coupled calculation, fast reactor, gfr 2400, nestle, neutronics, thermal-hydraulics, thermal feedback, validation. 1. introduction nowadays, the nuclear reactors are an appropriate option to decrease co2 production as the source with high energy capacity. however, the amount of fissionable 235u resources in the world are decreasing and, according to study [1], the entire conventional resources are sufficient for 240 years. together with public pressure to reduce long-lived radiological nuclides, there is necessity to introduce new alternative energy sources, which are capable to replace commercial thermal reactors in the future. therefore, the generation-iv international forum (gif) [2] identified perspective conceptual designs of fast reactor systems including the gas cooled fast reactor (hereinafter gfr 2400). however, multiple studies raise questions about the gfr 2400 technical feasibility [3] due to safety reasons. moreover, according to study [4] the recriticality issues may occur for fast reactor concepts and therefore the proper modelling of coupled neutronics with thermal-hydraulics is necessary to avoid these type of scenarios and for estimation of conceptual designs safety. the institute of nuclear and physical engineering (inpe) was involved in the gofastr project [5] that was focused on development of gfr 2400 design and inpe was also participating in allegro project which is oriented in research and development of gfr 2400 demonstration unit. the paper is focused on the coupled calculation of the gfr 2400 fuel assembly by the nestle code system. the nestle code is described in section 2 together with the validation of installation. section 3 describes the methodology of the coupled calculation and includes part about processing of appropriate parametric multigroup macroscopic cross-section library by scale code system, which is necessary for the calculation by the nestle code. the geometry and material properties of gfr 2400 are presented in section 4. the results of the scale and nestle calculations are shown in the last section 5 and are compared with the gfr 2400 reference study [6] performed during the gofastr project [5]. 2. nestle code 2.1. nestle code 5.2.1 description the main advantage of the nodal code nestle is the already implemented internally coupled neutronic and thermal-hydraulic iterative calculation path. the nestle code solves 2 or 4 group neutron diffusion equation utilizing the nodal expansion method [7]. −∇·dg∇φg + σtgφg = g∑ g′=1 σsgg′φg′ + χg k g∑ g′=1 νg′ σf g′φg′ (1) where dg is diffusion coefficient [cm], φg represents the neutron flux [cm−2·sec−1], σtg stands for total macroscopic cross-section [cm−1], σsgg′ for in-group 34 http://dx.doi.org/10.14311/app.2018.14.0034 http://ojs.cvut.cz/ojs/index.php/app vol. 14/2018 simulation of reactor fuel assembly with nestle code system scattering cross-section [cm−1], χg is the fission neutrons yield, k represents the multiplication factor (or fundamental eigenvalue), νg stands for average number of neutrons per fission and σf g represents the macroscopic fission cross-section [cm−1]. to calculate the above-mentioned diffusion equation, it is necessary to process appropriate parametric multigroup macroscopic cross-sections, according to the following equation: σ̂xg = a1xg + 2∑ n=1 a(n+1)xg (∆%c) n + a4xg ∆tc+ a5xg ∆(tfef f ) 1/2 + 3∑ n=1 a(n+5)xg (∆nsp) n (2) where σ̂xg represents the macroscopic cross-section for particular reaction type x and in the energy group g, ajxg are expansion coefficients of modified power series, ∆%c is the deviation from the reference coolant density [g·cm−3], ∆tc stands for deviation from the reference coolant temperature [ °f], ∆(tfef f )1/2 is the change in the square root of the effective fuel temperature in [ °f] from the reference condition and ∆nsp represents the deviation from the reference soluble poison concentration [cm−3·10−24] [7]. more details could be found also in the previous study [8]. 2.2. validation of installation the first version of the nestle code was developed at north carolina state university beginning in the late 1980’s and the latest version of the nestle code 5.2.1, which is available at nea data bank [9], was made in 2003. the license of newer developer version of the nestle code was not obtained by the inpe and therefore the whole calculation presented in the paper was performed on the version 5.2.1. the nestle code 5.2.1 was tested for sun computers with aix operating system based on powerpc (ppc) processor architecture with xlf fortran 77 compiler in 2004. therefore, the correct installation was difficult and the virtualization of ppc architecture (on ubuntu system 14.04.5 [10]) was necessary to perform for the successful nestle compilation by an updated version of xl fortran compiler (v. 15.1.5) [11]. the nestle calculation is a little bit inconvenient on virtualized ppc machine due to the more than 2 orders of magnitude higher computational time. based on the successful installation, the original code was slightly modified to comply with gfortran compiler, which is available on multiple architecture platforms. the results of the validation of installation by the benchmarks included in the installation package of the nestle code are shown in table 1 for the version compiled by gfortran compiler on ubuntu system 14.04.5 with amd64 architecture. the results shown in table 1 are identical for the platforms with different architecture. the difference in results of kef f is less than 1 pcm for almost all cases. the deviation of the calculated coolant and fuel temperatures from the benchmark values is insignificant (less than 10−1 °f). however, the calculated results of test case number 10 achieves higher deviation from the benchmark. the testcase was prepared for calculation of the boron acid concentration in the different fuel cycles of the core and obviously, at the end of the cycle (eoc), the boron concentration should be equal to 0. the calculated boron concentration in testcase number 10 is 1.7637 ppm (parts per million) and this difference should be discussed with the developer team in tennessee in the future. the next section describes the calculation process and the processing of macroscopic cross-section library for nestle code. 3. calculation methodology calculation scheme is presented in fig. 1. triton sequence of the scale code system [12] is used to process appropriate parametric homogenized multigroup macroscopic cross-section library for the nestle code. the modelled states of the nuclear system, referred as branches, are describing the deviation from the reference state of the nuclear system according to eq. 2 (i.e. difference in coolant density, coolant temperature and square root of effective fuel temperature for the case of gfr 2400). the above mentioned branches estimate the response of the nuclear system induced by the change of thermal-hydraulic parameters in the system. the triton sequence calls the deterministic multigroup code newt, which estimates the neutron flux in the investigated system and process homogenized collapsed and weighted macroscopic cross-sections according to the calculated flux. a simple script was developed at inpe to prepare appropriate approximations for nestle code according to the provided cross-section library made by the triton sequence. initially, the nestle code calculates the neutron flux distribution in the system. according to the calculated fission reaction rates, the thermal-hydraulic sub-channel calculation estimates the temperature distribution and density of the coolant in the system. according to these parameters, the macroscopic crosssection library is updated for next iteration step of the neutronic calculation within the nestle code. this process repeats until the calculation reaches the initial stopping criteria of the calculation. if the convergence of the solution is not achieved, it is necessary to update the input data and the conditions for both scale and nestle calculations. the next section describes the investigated case, which is the fuel assembly based on the gfr 2400 model. 35 f. osusky, s. cerba, j. luley et al. acta polytechnica ctu proceedings test case number calculated value of kef f benchmark value of kef f test case number calculated value of kef f benchmark value of kef f 1 1.000366 1.000364 14 1.000246 1.000245 2 1.000364 1.000364 15 0.989215 0.989215 3 1.000365 1.000356 16 1.000004 1.000000 4 0.999968 0.999968 17 1.006454 1.006454 5 1.000365 1.000363 18 1.009920 1.009920 6 1.122087 1.122004 19 1.011475 1.011475 7 1.121459 1.121373 20 0.977857 0.977858 8 1.121039 1.121117 21 1.005640 1.005640 9 1.200129 1.200129 22 0.999981 0.999982 10 0.999902 1.000088 23 0.999936 0.999936 11 0.979755 0.979755 24 0.999997 0.999997 12 0.989589 0.989589 25 0.999993 0.999992 13 1.088344 1.088343 26 0.999993 0.999994 table 1. validation of the nestle code installation compiled by gfortran compiler. figure 1. calculation iteration scheme of the simulation. 4. description of investigated case the current experience of the gas cooled reactor technology has more than 1000 reactor years of gas thermal reactor operation in europe [13]. moreover, slovakia operated the gas cooled thermal reactor a1 until mid-1980’s and is one of the possible countries to build the gfr 2400 demonstration unit (referred as allegro). therefore, the research institutes in slovakia are involved in the research and development of gas cooled fast reactor technology and the paper focuses on gfr 2400 concept. the power of the gfr 2400 reactor concept is 2400 mwth and the planned efficiency of the indirect brayton cycle is 45%. the used coolant is helium at 7 mpa pressure and the heat transfer is provided by three blowers with intermediate heat exchangers. the mass flow rate of the coolant inside of the fuel assemblies is 1213 kg·s−1 and the bypass flow rate of the coolant between the fuel assemblies is 60 kg·s−1. the high coolant volume fraction is characteristic for this concept (43 v/vtot%). the basic thermal-hydraulic parameters are shown in table 2 and these parameters are described in more detail in previous studies [8, 14] and are accompanied with the material composition. the cross sectional view of the gfr 2400 core is shown in fig. 2. the chosen fuel material is uranium and plutonium carbide (hereinafter u-puc) and one from the u-puc benefits are appropriate thermal properties for the cores with high temperature gradient, such as gfr 2400. the core is mixed from two zones with different enrichment: the inner core, with lower 36 vol. 14/2018 simulation of reactor fuel assembly with nestle code system parameter value parameter value thermal power [mw] 2400 primary coolant he primary pressure [mpa] 7 pressure drop in core [mpa] 0.143 mass flow rate [kg·s−1] 1213 bypass flow rate [kg·s−1] 60 core inlet temp. [ °c] 400 core outlet temp. [ °c] 780 table 2. basic thermal-hydraulic parameters [15]. figure 2. cross sectional view of the gfr 2400 core [15]. volumetric enrichment of puc 14.12%, and the outer core, with higher volumetric enrichment of puc of 17.65%. the purpose of two fuel zones with different enrichment is to achieve uniformly distributed neutron flux in radial direction of the core. however, the calculation will be performed with arithmetic average volumetric enrichment of plutonium isotopes 15.84% in the carbide fuel (table 3). the cross sectional view of the fuel pin is shown in fig. 3. the middle part in the fig. 3. represents the cut of u-puc fuel pellet. the gap between the fuel pellet and internal liner of cladding is filled with helium. the cladding is made from mixed structure of sicf /sic and this material is chosen thanks to its stable behaviour during high temperature gradients. however, the sic can be easily penetrated by fission gases and therefore, the special internal liner was introduced, made in the form of w14re. the w14re liner improves the fission gas retention of the fuel pin and moreover, prevents the carbonization of the cladding due to fuel-cladding interaction during the fuel cycle. the homogenization of the gap, internal liner and cladding with external liner was performed to achieve faster and stable convergence of deterministic calculation made by newt code during the processing of macroscopic cross-section library. the description of the homogenization process could be found in the fuel density [g·cm−3] 10.9002 %v in puc 15.84 %w in puc 15.81 %v in uc 84.16 %w in uc 84.19 %w pu-238 0.4065 %w pu-239 8.4314 %w pu-240 3.8995 %w pu-241 1.1142 %w pu-242 1.0991 %w am-241 0.1054 %w u-235 0.5770 %w u-238 79.5684 %w c-nat 4.7984 table 3. fuel material composition [15]. figure 3. cross sectional view of the gfr 2400 fuel pin [15]. 37 f. osusky, s. cerba, j. luley et al. acta polytechnica ctu proceedings homogenized cladding [g·cm−3] 2.9579 %w w-182 4.9824 %w w-183 2.6942 %w w-184 5.7989 %w w-186 5.4380 %w re-185 1.9861 %w re-187 3.3588 %w si-28 48.7308 %w si-29 2.5611 %w si-30 1.7492 %w he-04 0.0168 %w c-nat 22.6838 table 4. homogenized cladding material composition [16]. figure 4. cross sectional view of the gfr 2400 fuel assembly [15]. previous study [16] and the results of the fuel pin homogenization is shown in table 4. the gfr 2400 core consists of 516 fuel assemblies and the height of the core part is 1.65 m. each fuel assembly consists of 217 fuel pins and the cross sectional view of the fuel assembly is shown in fig. 4. 5. results 5.1. scale calculation as mentioned above, the triton sequence was used for processing of the parametric macroscopic crosssection library for the nestle code. the mesh generated for the calculation is shown in fig. 5. figure 5. mesh of the investigated case (the orthogonal mesh of the fuel pin is 4x4 and the mesh for whole fuel assembly is 40x40). the standard scale cross-section library, referred as v7-252, was used and is based on endf/b-vii.1 evaluated data. the 252 neutron group library was developed to adequately capture spectral and temperature effects important for reactor systems [12] and therefore was used for the branch calculation by the triton sequence. white boundary condition was considered during the calculation and the convergence criterion was set to the particular mixtures and for the eigenvalue (εkef f < 10−5). the calculations have converged for each branch calculation (inner and outer convergence was established). the reference conditions of nuclear system were the following: temperature 973.15 °c of the fuel; temperature 673.15 °c of the coolant; and the coolant density 4.9439 kg·m−3. the branch calculation was performed for the temperature range of the fuel until 1773.15 °c with 50 °c step. the coolant temperature range was up to 1473.15 °c with the same step (50 °c). the coolant density change corresponded to the coolant temperature change, which was mentioned above. the estimated multiplication factor was 1.21519 by the deterministic newt calculation for the reference case. the triton sequence successfully process the parametric homogenized multigroup macroscopic cross-section library in raw binary data-format. automatization script was used to prepare the approximations to the nestle code and the results are shown in next sub-section. 5.2. nestle calculation the boundary conditions of nestle calculation were set to reflective in radial direction and in axial direction to zero flux. the total flowrate of the coolant was 2.351 kg·s−1 for one fuel assembly and the reference power generation was 367.352 mw·m−3 for the fuel. the fuel assembly results of the nestle calculation are compared with results of the previous studies [8, 14] and with the reference gfr 2400 38 vol. 14/2018 simulation of reactor fuel assembly with nestle code system figure 6. results of the coolant axial temperature distribution. figure 7. results of axial power distribution. figure 8. results of the fuel axial temperature distribution. 39 f. osusky, s. cerba, j. luley et al. acta polytechnica ctu proceedings study [6]. in fig. 6, the results of axial coolant temperature distribution are compared between the nestle calculation, cfd calculation made by the flowvision code [8, 17] and between the gfr 2400 study [6]. the nestle calculation results show lower temperatures in the axial direction of the core than the results obtained by the cfd simulation and gfr 2400 study. in case of the cfd simulation and the gfr 2400 study, the modelled case was one pin with periodical/symmetrical boundary condition, which led to smaller cross-sectional area where the coolant flows (the cross sectional area is approximately 108.57 cm2) than in the case of one full fuel assembly (111.09 cm2). therefore, the effectiveness of the heat removal is higher in the nestle calculation case, what resulted in lower temperature at the outlet, compared to the cfd simulation and the gfr 2400 study. results of the axial power distribution are compared in fig. 7. the scale keno vi three-dimensional calculation was performed in the previous study [14]. the scale model was divided into 10 zones in axial direction and the model was externally coupled with cfd flowvision calculation. the standard scale cross-section library, referred as v7-238, was used in this case and is based on endf/b-vii.0 evaluated data. the investigated case of the scale-flowvision calculation was one fuel pin with reflective boundary condition in radial direction (symmetrical boundary condition in the case of cfd simulation) and zero flux boundary condition in axial direction (pressure outlet in the case of cfd simulation). neutronic material parameters of each zone were directly influenced by the cfd simulation. the axial reflectors below and above the core were not considered for the nestle calculation and for the scale keno vi calculation. therefore, the results of the simulated cases made by nestle and scale codes (fig. 7) are not consistent with gfr 2400 study [6], where the axial reflectors were considered. the application of the axial reflectors results in the higher fission reaction rate in the axial boundary regions than for the case without application of the axial reflectors. the results of axial fuel temperature are shown in fig. 8. and are reflecting the calculated axial power distribution. there is good agreement of the calculated cases by nestle and cfd simulations. an additional simulation was performed by the newt code, where the temperatures and densities were averaged, according to the performed nestle calculation. the result of the multiplication factor was 1.21401. the calculated multiplication factor of nestle code was 1.19661 and the difference from the newt simulation could have been caused by different approach of the deterministic-transport (typical for the newt code) and nodal-diffusion calculation (typical for the nestle code) codes. 6. conclusion the validation of the nestle code system installation is described in the paper. the results of the test cases were consistent with the benchmark results except one. in the test case number 10, the calculated boron acid concentration at the end of the fuel cycle was not equal to 0, what resulted in the bias of the multiplication factor for this test case. this behaviour will be discussed in the future with the developer team in tennessee. the coupled simulation of the gfr 2400 fuel assembly was performed by the nestle code and was compared with previous studies and with the reference gfr 2400 study from the gofastr project. the results made by the nestle code are in good agreement with cfd simulations made by the flowvision code. the axial reflectors above and below the core were not considered in the nestle simulation and this modelling approach resulted in the decrease of the power at the top and the bottom of the fuel assembly in comparison with gfr 2400 study, where these reflectors were included in the model. the technical feasibility of the gfr 2400 design raises questions and therefore the future activities with benchmarking are needed for this concept. the main reason is that the material composition of the design is quite unique and the appropriate benchmark does not exist for the gfr 2400 core. acknowledgements this work was financially supported by stu grant scheme for support of young researchers hopone no. 1323. references [1] nuclear energy agency and the international atomic energy agency. uranium 2016: resources, production and demand, 2016. https://www.oecd-nea.org/ndd/ pubs/2016/7301-uranium-2016.pdf. [2] t. abram. generation-iv nuclear power: a review of the state of the science. energy policy 36:4323–4330, 2008. [3] g. locatelli. generation iv nuclear reactors: current status and future prospects. energy policy 61:1503–1520, 2013. [4] h. ninokata. a study on recriticality characteristics of fast reactors in pursuit of recriticality-accident-free concepts. progress in nuclear energy 29:387–393, 1995. [5] european commission. gofastr : european gas cooled fast reactor. [2017-11-30], http: //cordis.europa.eu/project/rcn/96860_en.html. [6] european commission. gofastr : gfr2400 core neutronics and thermal-hydraulics characterization, 2012. work package 1.1: gfr conceptual design, http://www.janleenkloosterman.nl/reports/ gofastr_d1111_201203.pdf. [7] electric power research center, north carolina state university. nestle version 5.2.1: few-group neutron diffusion equation solver utilizing the nodal expansion method for eigenvalue, adjoint, fixed-source steady-state and transient problems, 2003. 40 https://www.oecd-nea.org/ndd/pubs/2016/7301-uranium-2016.pdf https://www.oecd-nea.org/ndd/pubs/2016/7301-uranium-2016.pdf http://cordis.europa.eu/project/rcn/96860_en.html http://cordis.europa.eu/project/rcn/96860_en.html http://www.janleenkloosterman.nl/reports/gofastr_d1111_201203.pdf http://www.janleenkloosterman.nl/reports/gofastr_d1111_201203.pdf vol. 14/2018 simulation of reactor fuel assembly with nestle code system [8] f. osusky. test case specifications for coupled neutronics-thermal hydraulics calculation of gas-cooled fast reactor. journal of physics: conference series 781(1):10, 2016. http: //iopscience.iop.org/1742-6596/781/1/012038. [9] nuclear energy agency. ccc-0641 nestle 5.2.1. [2017-11-30], http://www.oecd-nea.org/tools/ abstract/detail/ccc-0641/. [10] ubuntu. ubuntu-14.04.5-server-ppc64el. [2017-11-30], http: //cdimage.ubuntu.com/releases/14.04/release/. [11] ibm. getting started with xl fortran for little endian distributions. version 15.1.5. [12] oak ridge national laboratory. scale code system, 2016. version 6.2.1. [13] r. stainsby. gas cooled fast reactor research in europe. nuclear engineering and design 241:3481–3489, 2011. [14] f. osusky. definition of the thermal-hydraulic model of the gas-cooled fast reactor, 2017. in nuclear 2017: 10th international conference on sustainable development through nuclear research and education. pitesti, romania. may 24-26, 2017. pitesti: raten icn pitesti. issn 2066-2955. [15] z. perko. core neutronics characterization of the gfr2400 gas cooled fast reactor. progress in nuclear energy 83:460–481, 2015. [16] f. osusky. macroscopic cross-section processing for nestle code system, 2017. in apcom 2017 : proceedings of 23rd international conference on applied physics of condensed matter. strbske pleso, slovak republic, june 12-14, 2017. vol. 1. bratislava: publisher spektrum stu. isbn 978-80-227-4699-1. [17] capvidia. flow v ision: version 3.09.04, 2015. installation and setup guide. 41 http://iopscience.iop.org/1742-6596/781/1/012038 http://iopscience.iop.org/1742-6596/781/1/012038 http://www.oecd-nea.org/tools/abstract/detail/ccc-0641/ http://www.oecd-nea.org/tools/abstract/detail/ccc-0641/ http://cdimage.ubuntu.com/releases/14.04/release/ http://cdimage.ubuntu.com/releases/14.04/release/ acta polytechnica ctu proceedings 14:28–35, 2018 1 introduction 2 nestle code 2.1 nestle code 5.2.1 description 2.2 validation of installation 3 calculation methodology 4 description of investigated case 5 results 5.1 scale calculation 5.2 nestle calculation 6 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0091 acta polytechnica ctu proceedings 11:91–96, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app measurement methods of internal stress in continuous welded rail petr vnenk∗, bohumil culek department of transport structures, faculty of transport engineering, university of pardubice, studentska 95, pardubice, czech republic ∗ corresponding author: petr.vnenk@student.upce.cz abstract. this paper deals with the problem of internal rail stress estimation. it is based on a detailed research of contemporary situation in the field, presents basic outlines of the problem and sums up the major research areas and their possible applications in the current state of railway infrastructure management. the directions are divided into four categories in the paper: displacement methods, rail shifting methods, methods based on acoustoelastic effect and methods based on magnetoelastic effect. particular methods, both scientific and industrial, are presented in their section respectively. every method that is presented within the scope of this paper is briefly described and its advantages and disadvantages are mentioned. in the end, potential of application of some of the presented methods in the practical use is discussed. keywords: rail, stress, measurement, cwr, temperature. 1. introduction continuous welded rail, abbreviated as cwr, is a part of railway superstructure that has been under continuous development since its beginnings in the first half of the twentieth century. the implementation of continuous welded rail caused a major qualitative improvement in the railway transport. one of the biggest challenges of this improvement was the mitigation of the thermal stress effect. at the moment of welding, there is a certain stress state which equals a temperature called "neutral temperature". the neutral temperature is characterized by the attribute that when a welded rail reaches the neutral temperature, the rail is free from thermal stress. when the rail temperature is higher than the neutral temperature, compressive stress in rail emerges. at extreme temperatures, the compressive stress induced in the rail can cause railway track buckling, also known as "sun kink". on the contrary, when the rail temperature is lower than the neutral temperature, tensile stress in rail emerges. at extreme temperatures, the tensile stress induced in the rail can cause the rail to be pulled apart. the thermal stress in the continuous welded rail is expressed as σ = −α · ∆t · e (1) where α = 12 · 10−6k−1 (2) and e = 210 · 109pa (3) in the czech republic, the permitted welding temperature for rails ranges from 17 °c to 23 °c [1]. considering the minimum rail temperature at -30 °c and the maximum rail temperature at 60 °c (for rails situated in the sunny locations and thus heated), the difference between the extreme temperature and the neutral temperature can reach as many as 53 °c [2]. assuming such conditions, the thermal stress in continuous welded rail is σ = 133.56 · 106pa (4) this stress as such is not high enough to cause failure of continuous welded rail. however, there are other contributing factors, too. the thermal stress is cyclic, repeating in both daily and yearly cycles. trains cause further heavy dynamic loading which is cyclic, too. all these loadings are countable. it appears, however, that the neutral temperature is not constant over a period of time. based on the location of the rail, it seems the neutral temperature is shifting. this shift of the neutral temperature can reach even several degrees of celsius. this difference has to be added to the ∆t parameter, eventually increasing the thermal rail stress. combinations of the above mentioned stresses and their cyclic character have adverse influence on continuous welded rails. unlike the other items, it is very difficult to predict the shift of the neutral temperature. this article presents contemporary findings in 91 http://dx.doi.org/10.14311/app.2017.11.0091 http://ojs.cvut.cz/ojs/index.php/app petr vnenk, bohumil culek acta polytechnica ctu proceedings this area and suggests possible solutions to the evaluation of the thermal stress in rail with the influence of the neutral temperature shift. 2. displacement methods displacement methods of cwr stress measurement work on the simple mechanical assumption of change of a relative distance between two points of material, rail in this case. in newton’s mechanics, relatively stiff rail behaves according to hook’s law if unfastened. however, even the fastening systems ensure only a certain level of rigidity and, provided the axial force in rail is strong enough, a relative displacement can be observed even at fastened rail sections. as of now, some solutions using displacement are available as presented below. 2.1. research of the gdansk university of technology this research was carried out in the mid-seventies at the gdansk university of technology, poland. a displacement sensor with accuracy of 0.001 mm has been applied for a measurement on a 200 mm long rail section. next to the cwr measurement spot, a reference rail section has been placed. this reference rail section has been set free to lengthen and shorten. it was able to observe soon that the stress varies over different sections of the rail even if the temperature of the rail was the same at measured locations. it is believed that this effect is caused by slip of rail in the fastening points. such a slip can be derived by several sources like tangential forces of traffic, but also by exposure of some parts of rail to sun and location of some parts of rail to completely shady place. such a different exposure to sun results in different daily mean temperature and, subsequently, to a stress gradient in the rail. this revealing is a good asset of the measurement. the method itself, however, is not suitable for commercial use as it is very costly, time demanding and requires a reference rail to be placed at the measurement spot. nevertheless, it can well serve as a reference method for calibration of other methods [3]. 2.2. the calibrated length of rail method this method, which has supposedly been applied in china since 1982, consists of length comparison of a freely placed 50 m long steel strip with a selected rail section originally of the same length. the principle of this method, however, appears to be the same like in the research of the gdansk university of technology, only modified for a bigger span. author’s daring claim that this method is the only one in the world without "fatal" shortcomings and can be operated widely does not match the state of the research in this area after the publication date, because much of research has been emerging in this field [4]. 2.3. measurement with bi-directional fbg strain sensors one of the most recent papers covering this topic, from 2015, presents a method of displacement measurement in a rail web in longitudinal and vertical direction using optic fbg sensors. in general, this measurement resembles strain gauge measurement. however, the advantage of fbg sensors is the possibility to measure even temperature, apart from mere strain. moreover, it is characterized by its longer durability in relation to strain gauge sensors. provided the sensor is attached to the rail web prior to cwr installation (i.e. at rail neutral temperature or known rail tension), it keeps measuring the deformation as long as the sensor works. this can be theoretically reached using strain gauges only, although the durability appears as an important factor here. [5]. 2.4. vortok measure and detect another option to measure displacement can be reached by using the measure and detect sensor produced by vortok ltd. company. simple sensor that is able to measure variety of parameters is made ready to be screwed into a drilled hole in a rail web. the manufacturer does not share any detailed information of what principles this device uses. it looks, however, like a piezoelectric sensor with accessories that is incorporated into a screw-in dowel. this robust shape most probably predicts a displacement method with a long durability. a disadvantage of this method is a necessity of calibration. after calibration is made, either by fastening of the device on a stress-free rail or by using another calibration method, the value of internal stress in rail can be obtained in a stable manner. the options to set up automatic data transmission via email or wi-fi even from a moving train make this sensor a user-friendly device [6]. 3. rail shifting methods rail shifting methods of cwr stress measurement use the principle of bending stiffness dependence on axial tension in rail. provided the rail is in tension, the higher force is necessary to be applied to laterally shift the rail the higher the axial force in the rail is. apparently, these methods are viable only when the rail is in tension, otherwise failure by buckling may appear. 3.1. verse method this method has been presented by vortok ltd. company from the united kingdom. in the first place, the rail shall be released from all fastenings in the length of 30 m. afterwards, the central part of the unfastened section of the rail shall be lifted into a certain height. the force that is applied to move the rail is related to the axial force in the rail. wide use of this method is restricted by some disadvantages. the most obvious one is the limitation to tensile stress in the measured rail. the operator needs 92 vol. 11/2017 measurement methods of internal stress in cwr to be sure that there is a tensile stress as problems with re-installation of the rail into the original position could occur, leave apart the danger of operator’s injury or material damage. another disadvantage is the need to unfasten the rail in the length of 30 m. this operation requires additional time and subsequently prolongs track closure. moreover, in case the measured section is situated in radius, re-installation of the rail can be difficult, too, as the condition of tensile stress in rail is inevitable. even though the disadvantages are important, this method appears to be one of the most widely spread to measure the cwr stress nowadays. the manufacturer produces their own device for rail lifting and stress measurement. light aluminium design of this device enables easy transport, manipulation and operation [7]. 3.2. use of tamping machine in a paper published in 2009, scientists at gdansk university of technology look into a possibility to use tamping machine for stress measurement in cwr. tamping machine is capable to shift rails in both vertical and horizontal direction when adjusting railway track geometry. the goal of this research was to use this movement to estimate the stress in cwr. the principle of this measurement is similar like at the verse method, including the disadvantages. however, the disturbing signal produced by the operation of tamping machine turned out to be too high and the attempts to get some cwr stress data were unsuccessful [8]. 4. methods based on acoustoelastic effect the acoustoelastic effect is a physical effect of change of sound velocity based on change of mechanical stress in elastic continuum. for measurements, ultrasound is used. the definition, however, can cover vibration, as mechanical waves, too. in such a case, vibrometer is used as a measurement tool. it is always necessary to execute a calibration measurement on the site with known temperature and mechanical stress as the sound velocity in rail is not dependent only on mechanical stress, but on steel microstructure, too. this calibration measurement can be performed either prior to cwr installation, or once the cwr is installed. however, in the latter only when the rail is cut, unfastened and welded again. as a consequence, similar disadvantage like at displacement methods emerges, which means the need to start the measurement at the neutral temperature. the advantage, on the contrary, is in possible high durability of ultrasound sensors and their attachment to measurement place. 4.1. j. szelazek measurement j. szelazek has worked out two ultrasound measurement procedures in the nineties of the twentieth century. in the first procedure, ultrasound signal is emitted vertically from the top of rail head to the bottom of rail foot where it reflects off and moves back to the receiver, which can be the same device as the emitter. the second procedure uses separate emitter and receiver and those are placed to the side of rail head horizontally next to each other. the signal moves through the rail head lengthwise, in this case. the first procedure showed as inapplicable, as the signal dispersed due to uneven rail head profile. the major cause of this unevenness is the rail head wear. placing the signal receiver to the bottom of rail foot did not improved the results and, moreover, the signal path was reduced to one half. the second procedure appeared more promising, but even at this measurement, disturbances appeared and the results were influenced by strong dispersion [9]. 4.2. research of the university of florence in 2007, italian scientists worked out the second procedure of j. szelazek. this procedure has been tested for 2 years and appears viable. however, it shares all the common disadvantages of ultrasound methods and the necessity to perform calibration measurement in situ [10]. 4.3. measurement of stresses using the polarization of rayleigh surface waves two us researches work with measurement of polarization of rayleigh surface waves. measurement of dependence of the rayleigh surface waves polarization on change of mechanical rail stress turned out not just to be more robust than measurement of dependence of the rayleigh surface waves velocity on change of mechanical rail stress, but also to be easier to detect. michael d. a. junge’s work deals with general measurement of mechanical stress in material based on rayleigh surface waves [11]. stefan hurlebaus’ report deals with application of this procedure on rail stress research. on page 42, the author presents that the polarization of rayleigh surface waves is dependent on the value of axial force in rail and that a further research is recommended [12]. 4.4. vibration measurement two publications, a diploma thesis and a research report, deal with vibration measurement of the internal rail stress. both of them are outputs of one research. the results of the rail stress determination from wave lengths of vibrating sensor seem well in laboratory conditions when a steel bar or a new rail is used. in case of worn rail, and this is very important for practical application, it is possible to get certain results, but these results are not that clear like in the previous cases. when performing a field measurement, the results of measurement on worn rail are much worse. 93 petr vnenk, bohumil culek acta polytechnica ctu proceedings as of yet, this method is inapplicable. in case it is possible to diminish the problems, this method can have a good perspective [13, 14]. 5. methods based on magnetoelastic effect the magnetoelastic effect is inverse magnetostrictive effect. it is presented as a change of magnetic susceptibility in relation to change of mechanical stress in material. the magnetoelastic effect is also known as the villari effect. 5.1. measurement of magnetic hysteresis in the nineties, d. utrata tested the possibility of application of the magnetoelastic effect on the rail stress measurement. he measured the dependence of coercivity and remanence of the magnetic field in rail on different surface conditions, like scaled rail surface, or milled rail surface. however, the comparison of the obtained data has shown various results based on the input type of rail. surprisingly, the only data that match were on the milled running surface and scaled base underside. d. utrata assumes, in the end, that either a calibration measurement or a new approach to data interpretation has to be delivered in order to make this method applicable. a. wegner needed 400 calibration sensors to run a measurement of 3 meters long rail using the hysteresis method, which does not contribute to viability of this method [14, 15]. 5.2. measurement of magnetic barkhausen effect japanese scientists tsuchima and enokinozo have studied the reaction of the barkhausen noise on mechanical stress in steel plate. it is possible to get some relation from the graphs of the mutual dependence. however, the authors found it difficult to estimate the influence of time and get a value of the mechanical stress. additionally, they suggest application of the chaos theory to get a better interpretation of the results [16]. measurement using the barkhausen noise together with measurement of magnetic permeability is used by elektro-thermit gmbh & co. kg. company to determine the internal rail stress. their method requires calibration on a rail test sample. their product has been awarded a certificate of the railway infrastructure manager of denmark, banedanmark. on the contrary, the manufacturer appears not interested in sale of their product and only offers to perform a measurement in the required section by themselves [3, 17]. 5.3. metal magnetic memory measurement collective of authors from the university of nanjing have carried out an experiment of measuring the effect called metal magnetic memory on a steel test sample. figure 1. dhp(y)/dx for applied tensile stress on a steel sample [18]; see the linear dependence in the elastic region. contrary to the barkhausen noise, this effect does not require external magnetization of the sample. the effect is activated anywhere in the natural magnetic field of the earth. when the loading increases, the rotation of the weiss magnetic domains in metals raises the value of the magnetic flux around the central point and this increases the gradient of the magnetic flux which is in elastic region linearly dependent on the stress in the sample. graph of the linear dependence of the magnetic flux gradient on the mechanical stress in a steel sample as measured by wang, et. al. [18] is presented in figure 1. the linear dependence in the elastic region is completely sufficient, because rail geometry deformation emerges earlier than plastic deformation in the axial direction of the rail [18]. 6. discussion every presented method has its limitation. displacement methods can be related to widely spread strain gauge measurement. this has been used in many industrial application; in this case, however, the biggest problem seems to be the necessity to measure an absolute value of stress, not a relative one, which is the common purpose of strain gauge measurement in general. rail shifting methods, especially the verse method, seem to be one of the most reliable at this time. the reliability is underpinned by its simplicity. however, more extensive use of this method is very limited through many disadvantages it inevitably drags along. movement of a rail is a complicated operation, which needs expensive tools and is very time demanding. re-installation of the rail requires verification that the rail is in good condition after the measurement and safety of operation on the track is kept. additionally, performing such a measurement in radius is much more complicated and performing such a measurement when the rail is under a compressive stress is excluded. considering the compressive stress 94 vol. 11/2017 measurement methods of internal stress in cwr in rail is usually during the day, the limitations of use of this method are vast. methods based on the acoustoelastic and magnetoelastic effect have the potential to become viable in the non-destructive rail stress measurements. as of yet, however, the problems of getting clear values of stress were not mitigated and residual stresses, difficult conditions of field measurement, or even various conditions of the rail itself (wear, rust, etc.) prevented the research to reach a satisfactory solution. 6.1. further research orientation out of the contemporary possibilities, there are some ways to be worked out in order to either get satisfactory results or to discover serious obstacles that hinder their application. (1.) application of the metal magnetic memory method on internal rail stress measurement seem like a research that has not been done yet. as mentioned in section 5, this method does not require external magnetization [18] and this is a great asset for possible further application. if the measurement of the gradient of magnetic flux in rail results in a similar dependence as shown in figure 1, the potential of applicability of this method would be great [18]. (2.) vibration measurement, or modal analysis could also offer an approach to obtain data of internal stress in cwr. however, damping caused by rail fastenings could have too high influence on the results. as presented in section 4, the vibration analysis performed by the team of the university of illinois at urbana-champaign did not produce any applicable results even after they released some fastenings, which is a procedure that is disadvantageous as such [13, 14]. (3.) provided the stress could be evaluated from a surface layer of rail only, i. e. there was a mutual relation between surface stress and stress in the core of the rail, another possible method emerges. method of dependence between passive layers formation in electrolyte and mechanical stress in the surface layers has been shown in an experiment of stainless steel in a normal sulphuric acid bath [19]. (4.) another option is to work out the measurement of stresses using the polarization of rayleigh surface waves. the papers presented in section 4 suggest a further research can have a potential to find out a viable solution [11, 12]. (5.) finally, if strain gauge is able to measure a quasi static load change in a long term, a statistical evaluation of measurements could provide a range of data that could be sufficient for prediction of neutral temperature level and subsequently stress level in rail. if such a research is biased towards a certain limited location, an applicable methodology can be a reasonable output. 7. conclusions many attempts to measure the neutral temperature or stress in continuous welded rail have been undertaken. displacement and rail shifting methods like vortok measure and detect or verse have been finalized in an applicable product and offered on market. these methods come out of mechanical solutions. however, these methods suffer from many limitations and requirements of special conditions that are necessary to be created in order to use the methods. later in time, non-destructive methods based on the acoustoelastic or magnetoelastic effects have been tested. as of yet, none of them reached the level of being reliable enough to compete with the above mentioned ones. nevertheless, there are some promising results and further research in this area, as described in section 6, could prove viability of some of them. provided strain gauges are capable of measuring quasi static load in a long term and a thorough strain gauge measurement is done, the value of stress in a continuous welded rail could be estimated on a certain geographical area even from statistical data measured by strain gauges. finally, the possible solution can lurk in another method of stress evaluation in a continuum, a method which is not yet known to the authors of this paper. acknowledgements the work was created in connection with the scientific research project of the university of pardubice no. sgs_2017_009. the authors are grateful for the support. references [1] sprava zeleznicni dopravni cesty, s.o. szdc s 3/2 bezstykova kolej, 2013. kapitola iii, cast a, bod 110. [2] sprava zeleznicni dopravni cesty, s.o. szdc s 3 dil iv zeleznicni svrsek. kolejnice ve zneni zmen c. 1 a 2, 2014. kapitola iv, bod 43. [3] w. koc, et. al. okreslanie wartosci sil podluznych w szynach toru bezstykowego. problemy kolejnictwa 163, 2014. [4] l. xing-han. the "calibrated length of rail" method for measuring rail temperature stress. arema proceedings of the 2000 annual conference 2000. [5] p. wang, et. al. longitudinal force measurement in continuous welded rail with bi-directional fbg strain sensors. smart materials and structures 25, 2015. [6] vortok, ltd. vortok measure and detect, 2012. vortok, [2016-06-07], http://www.vortok.com/rail-stress-management/ vortok-measure-and-detect-2. [7] vortok, ltd. verse, 2012. vortok, [2016-05-31], http: //www.vortok.com/rail-stress-management/verse. [8] w. koc, a. wilk. investigations of methods to measure longitudinal forces in continuous welded rail tracks using the tamping machine. journal of rail and rapid transit 223(1), 2009. 95 http://www.vortok.com/rail-stress-management/vortok-measure-and-detect-2 http://www.vortok.com/rail-stress-management/vortok-measure-and-detect-2 http://www.vortok.com/rail-stress-management/verse http://www.vortok.com/rail-stress-management/verse petr vnenk, bohumil culek acta polytechnica ctu proceedings [9] j. szelazek. monitoring of thermal stresses in continuously welded rails with ultrasonic technique, 1998. ndtnet, 1. june 1998, [2016-05-30], http://www. ndt.net/article/dresd97/szelazek/szelazek.htm. [10] d. vangi, a. virga. a practical application of ultrasonic thermal stress monitoring in continuous welded rail. experimental mechanics 47, 2007. [11] m. d. a. junge. measurement of applied stresses using the polarization of rayleigh surface waves, 2003. atlanta, georgia, school of civil and environmental engineering, institute of technology. [12] s. hurlebaus. determination of longitudinal stress in rails, 2011. college station, texas, transportation research board of the national academies. [13] r. weaver. vibration measurement of rail stress, 2006. champaign, illinois, transportation research board of the national academies. [14] m. j. koob. the development of a vibration technique for estimation of neutral temperature in continuously welded railroad rail, 2005. champaign, illinois, university of illinois at urbana-champaign. [15] d. utrata, a. strom, m. negley. stress measurement in railroad rail using ultrasonic and magnetic technique. review of progress in quantitative nondestructive evaluation 14, 1995. [16] y. tsuchida, m. enokinozo. stress evaluation by chaotic characteristics of barkhausen noise. review of quantitative nondestructive evaluation 22, 2003. [17] a. wegner. stress-free temperature monitoring using different measuring technologies – experiences and assessment. proceedings of 10th international heavy haul association conference 2013. [18] p. wang, et. al. stress measurement using magnetic barkhausen noise and metal magnetic memory testing. 17th world conference on nondestructive testing 2008. [19] f. navai. effects of tensile and compressive stresses on the passive layers formed on a type 302 stainless steel in a normal sulphuric acid bath. journal of materials science 30, 1995. 96 http://www.ndt.net/article/dresd97/szelazek/szelazek.htm http://www.ndt.net/article/dresd97/szelazek/szelazek.htm acta polytechnica ctu proceedings 11:91–96, 2017 1 introduction 2 displacement methods 2.1 research of the gdansk university of technology 2.2 the calibrated length of rail method 2.3 measurement with bi-directional fbg strain sensors 2.4 vortok measure and detect 3 rail shifting methods 3.1 verse method 3.2 use of tamping machine 4 methods based on acoustoelastic effect 4.1 j. szelazek measurement 4.2 research of the university of florence 4.3 measurement of stresses using the polarization of rayleigh surface waves 4.4 vibration measurement 5 methods based on magnetoelastic effect 5.1 measurement of magnetic hysteresis 5.2 measurement of magnetic barkhausen effect 5.3 metal magnetic memory measurement 6 discussion 6.1 further research orientation 7 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0114 acta polytechnica ctu proceedings 15:114–119, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app relation between stationary and moving pedestrian load models jan štěpánek∗, jiří máca czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jan.stepanek@fsv.cvut.cz abstract. the dynamic analysis of footbridges became common in recent years, but the approach to modelling pedestrian induced load is still inconsistent. this paper compares two different ways of modelling single pedestrian load. the response of the structure to stationary pulsating force is compared to the results given by the same force moving along simply supported beam. 300 randomly generated beams were used in this parametric study to describe the relation among these two approaches and parameters of the structure. the proposed formula was successfully applied to get more realistic estimation of response of the structure to a single walking pedestrian using stationary pulsating force. keywords: footbridges, pedestrian load models, forced vibration, response of structure. 1. introduction footbridges designed and built during recent years are very slender and flexible structures in many cases. such footbridges may be very susceptible to pedestrian induced load which may cause an uncomfortable level of vibration. natural frequencies of such structures are often very close to walking frequency of pedestrians, and therefore the resonance effect may occur. a number of studies have been made in past decades to define the procedure of effective and safe design of footbridges. a footbridge designed this way has to satisfy the criteria of serviceability limit state (sls). to be able to detect the risk of unacceptable vibration during design phase the dynamic analysis is often necessary. the european codes define the maximal level of acceleration of a bridge deck which can be reached to keep the sufficient comfort of pedestrians. numerical simulation can reveal the sensitivity of the structure due to pedestrians but unfortunately there is still a lack of pedestrian load models in current eurocodes. many single pedestrian load models have been developed [1, 2] but most of them assume the numerical analysis using moving forces or moving biodynamic models. nevertheless, most of the commercial programs used for static design of bridges do not provide the option of calculation with force moving along the structure. stationary harmonic pulsating force was presented in guidebook [3] to simulate single walking pedestrian. the advantage of this estimation is its simplicity when it comes to numeric analysis in static software. the main disadvantage of this load model is that it does not consider the time that one spends on the bridge. the stationary pulsating force without time limitation may cause overestimation of the response of the structure which leads to unnecessary design changes and potential increase of cost of the structure. this paper focuses on comparison of approaches of modelling single pedestrian load by stationary and moving pulsating force. the goal is to find a simply applicable formula which combines these two approaches. the formula can be used to find realistic estimation of response of a simply supported beam using stationary pulsating force. 2. pedestrian load models pedestrian load can be divided into several categories. normal traffic on the bridge can be defined as a single pedestrian load, group of walking people or stream of pedestrians. exceptional load situations (jumping, running, kneeling or vandalism) must not excite the intolerable vibrations either [2]. to be able to evaluate the response of the structure in normal operation, it is necessary to define a simple but accurate dynamic single pedestrian load model. 2.1. single pedestrian load dynamic force produced by a single person can be represented by three components in perpendicular directions: vertical, horizontal lateral and horizontal longitudinal components [1]. horizontal longitudinal force is usually considered negligible. the lateral component is caused by moving centre of gravity horizontally while walking. the frequency of such force is half than the stride frequency. a vertical component with the stride frequency is the cause of most of the documented problems with uncomfortable vibrations of footbridges. only the vertical component is further investigated in this paper. vertical pedestrian load, assuming to be perfectly periodical force, can be formulated by fourier series in equation (1). numerous different approximations of such formulation were developed in the past years [1, 4]. 114 http://dx.doi.org/10.14311/app.2018.15.0114 http://ojs.cvut.cz/ojs/index.php/app vol. 15/2018 relation between stationary and moving pedestrian load models fv (t) = g + g n∑ i=1 αi sin(2πifpt + ϕi) (1) where: fv (t) vertical pedestrian load g load corresponding to pedestrian’s weight αi fourier coefficient of the respective member fp step frequency ϕi phase shift n total number of included members various coefficients αi and ϕi can be found in literature [1, 4, 5]. in general, it can be said that including more than three or four harmonic members has no effect on the final load due to low values of αi for higher components. the worst-case scenario, when the stride frequency is equal to one of the natural frequencies of the bridge, is usually adopted for simulating a single pedestrian walking along the structure [1, 4, 6]. pedestrian walking velocity vs also plays an important role in serviceability verification of the structure. the speed vs can be taken by eq. (2) vs = fsls (2) where fs is step frequency and ls is step length. the recommended mean value for ls also varies in different studies. guidebook [4] recommends ls = 0.9 m. according to guidebook [3] ls is dependent on step frequency and for 2 hz (defined as normal walk) the typical step length ls = 0.75 m is considered. 2.2. stationary pedestrian load model the stationary pulsating force can be a simple way to find out the response of structure to dynamic pedestrian load. the guidebook [6] recommends to apply stationary harmonic force with amplitude 280 n to the most adverse position of the structure. the value of the force is taken as 700×0.4=280 n which is in accordance with average pedestrian weight 70 kg and the first fourier coefficient 0.4. other approaches take the movement of a person in account. the amplitude of applied force is multiplied by the reduction constant r. allen and murray give r=0.7 while grundman takes r=0.6 [1]. the formula for reasonable reduction factor r for simply supported beam is set in further parametric study. 2.3. moving pedestrian load model the moving load models can be used to define the response of the structure to pedestrian load in a very realistic way [5, 7]. examples of more sophisticated load models, including biodynamic models and inverted pendulum can be found in [8]. many force models are summarized in [1, 4] and most of them contains first fourier coefficient α1 ≈ 0.4. for this reason, one of the simplest models proposed by charles and hooprah was used in this study to simulate a walking pedestrian. this model consists only of static force g and first harmonic member for α1 = 0.4. this simplified approach is also tolerated in guidebook [2] for simulation of a walking pedestrian. 3. parametric study in this chapter a numerical analysis of 300 simply supported beams with randomly generated parameters was performed to find the formula for the reduction factor r. both stationary a moving pulsating force were used to find the peak acceleration in vertical direction. value of r defined in eq. (3) was taken for every single structure. finally, the function of r was found using the least-squares approximation with parameters of the tested beams used as the basis functions. rcal = amov ast (3) where amov is peak acceleration all over the structure calculated using moving force model and ast is acceleration induced by stationary force model without time limitation. 3.1. set of simply supported beams all the tested structures share the scheme of simply supported beam. every beam was generated so that one of the three lowest natural frequencies equals to 2 hz. the i-th natural frequency of simply supported beam can be calculated from analytical equation (4). ωi = π2i2 √ eiy µl4 (4) where: ωi i-th natural circular frequency e modulus of elasticity iy cross sectional second moment of area µ mass of one meter of length l span length iy can be expressed: iy = µl4ω2i π4i4e (5) and µ is defined in eq. (6). µ = ρa + µad (6) if ωi = 2π × 2 hz, every other parameter excluding iy can be randomly generated and iy is found using eq. (5). this procedure allows to generate simply supported beams with i-th natural frequency equals to 2 hz. assuming all the theoretical structures have rectangular cross-section, its width b and height h can be found by solving system of two equations (7) and (8). a = bh (7) 115 jan štěpánek, jiří máca acta polytechnica ctu proceedings iy = bh3 12 (8) where: a cross sectional area ρ material density µad mass added on each meter of length without static effect a single material – reinforced concrete c 30/37 was chosen for all the theoretical beams. the ranges of randomly generated parameters: ρ = 2500 kg/m3 standard density of concrete c 30/37 e = 32 gpa modulus of elasticity of concrete c30/37 ωi = 12.5664 circular frequency for fi = 2 hz. l = 10-50 m the typical span length of footbridges with static scheme of simply supported beam. a = 0.03-3 m2 wide range of cross sectional area which allows sufficient diversity of structures in the set. µad = 0 300 kg/m i = 1-3 one of the first three natural frequencies equals to 2 hz ξ = 0.008-0.02 damping ratio of concrete footbridges [1] the goal of this chapter is to assemble a set of simply supported beams with wide range of parameters so that the statistical analysis can be done using the leas-squares method. it means that some of the structures could exist just as mathematical model, not a real structure, because of their slenderness or high susceptibility to vibration. some other conditions of the properties of generated beams are set to reduce the amount of such structures in the list: • h > 0.1 m, b > 0.1 m minimal height and width of the cross section • 0.1 < h/b < 10 minimal and maximal ratio of height and width • h/l < 0.1 maximal ratio of cross sectional height and span length • b/l < 0.5 maximal ratio of width and span length the final set of the structures is composed of three parts. each part contains 100 beams with i-th natural frequency 2 hz for i = 1-3. every structure is modelled using finite element method (fem) in matlab and consists of 50 beam elements. the shear deformation of elements is neglected according to bernoulli-navier hypothesis. this simplification can be applied thanks to condition h/l<0.1 . 3.2. response of the structure 3.2.1. response to stationary pulsating force the i-th mode shape with corresponding natural frequency 2 hz must be found at first. the vertical harmonic force fv (t) with amplitude of 280 n and frequency fv = 2 hz (9) is then applied in the most adverse position of the bridge. fv (t) = 280 sin(2πfvt) n (9) maximal displacement of the structure is found using newmark-β method for solving equation of motion: kr(t) + cṙ(t) + mr̈(t) = p(t) (10) where: k stiffness matrix c damping matrix m mass matrix r(t) vector of displacement ṙ(t) vector of velocity r̈(t) vector of acceleration p(t) load vector created from fv (t) maximal acceleration can also be found solving the system of 2n linear equations (11) for system of n degrees of freedom. this approach is technically much easier but it does not provide any information about time history. [ (k − ω2m) −ωc ωc (k − ω2m) ]{ rs rc } = { fs fc } (11) where ω is the circular frequency of pacing force and r(t) and p(t) are expressed by two components: r(t) = rs sin(ωt) + rc cos(ωt) (12) p(t) = ps sin(ωt) + pc cos(ωt) (13) the amplitude of acceleration r̈ of such harmonic motion is given by eq. (14). r̈ = ω2 √ r2s + r2c (14) the reference value ast is finally found as maximal value of acceleration vector r̈ in vertical direction. 3.2.2. response to moving pulsating force the pedestrian load model, used in this section, was adopted from [1]. it consists only of static weight of pedestrian g = 70 kg and the first member of fourier series α1 = 0.4 and is expressed by eq. (15). the model is moving along the structure. to define the velocity of its movement, the step length needs to be defined. the step length ls = 0.75 m which is appropriate for step frequency was adopted from [3]. the walking speed was set according to eq. (2). 116 vol. 15/2018 relation between stationary and moving pedestrian load models figure 1. separate contributions of basis functions. fv (t) = 700 + 280 sin(2πfvt) n (15) the peak acceleration amov in vertical direction reached on each structure was analysed using newmark-β method. 3.3. evaluation the peak accelerations ast induced by stationary force model and amov induced by the moving force model were evaluated for all the structures which resulted in 300 values of reduction coefficient rcal eq. (3) the least-squares method was used for determination of the relation between reduction coefficient r and the parameters of the structure. eleven basis functions bj were defined: (1.) l span length (2.) µ mass for 1 meter of the beam (3.) iy cross sectional second moment of area (4.) µl total mass of the structure (5.) 1 l multiplicative inverse of span length (6.) 1 µ multiplicative inverse of mass for 1 meter of the beam (7.) 1 iy multiplicative inverse of cross sectional second moment of area (8.) i the order of frequency which equals to 2 hz. (9.) constant (10.) ξ1 damping ratio of the first mode shape (11.) 1 ξ1 multiplicative inverse of damping ratio of the first frequency the coefficients βj of basis functions were found solving system of equations (16).   x11 x12 · · · x1n x21 x22 · · · x2n · · · · · · ... ... xm1 xm2 · · · xmn     β1 β2 ... βn   =   y1 y2 ... ym   (16) where xij = 300∑ k=1 bi,kbj,k and yi = 300∑ k=1 bi,krcal,k and k stands for number of the structure in the list. the first calculation was primarily performed in order to define the parameters which have the greatest influence on the value of reduction factor. four functions (no. 5,8,9 and 10) were chosen for further analysis. the second calculation was performed using only four selected functions and the respective coefficients β were found. the function rcal can be approximated by rapp which is expressed as a linear combination of basis functions (17). the figure 1 shows the contributions of chosen basis functions. rcal ≈ rapp = ∑ bjβj (17) the formula for r is finally set by equation: r = −6.248 l + 0.0495i + 14.42ξ1 + 0.726 (18) and is usable in reduction of the stationary pulsating force model of single walking pedestrian according to eq. (19). fv (t) = r280 sin(2πfvt) n (19) as can be seen in figure 2, all the values of rcal falls in range between 0.5 1 thus the range of meaningful values is defined as: 0.5 ≤ r ≤ 1 (20) 117 jan štěpánek, jiří máca acta polytechnica ctu proceedings figure 2. calculated and approximated values of r. the average relative error of the approximation ∆r calculated according to eq.(21) is equal to 2.3 %. taking all the uncertainties of pedestrian load models into account, the value of relative error is sufficiently small. ∆r = ∣∣∣∣rcal − rapprcal ∣∣∣∣ (21) 4. verification of the proposed formula verification of the proposed formula has been performed on numerical model of experimental structure presented in [5]. the experimental footbridge is prestressed concrete beam with t cross section and short overhangs (figure 5). the tuned mass damper (tmd) was installed under the midspan of the beam to show its influence on the susceptibility to vibration. tmd allows the options to be locked, which is performed by stiff connection to the structure. the moving load model used in section 3 provided very accurate results of response of the structure in comparison to load tests [5]. calculations were performed with both locked and free tmd to verify the formula for reduction factor r (18). the first natural frequency of the structure f1 = 1.98 hz, and therefore the step length of 0.75 m can be applied according to guidebook [3]. several parameters of the structure are required to estimate the reduction factor r: • l = 19 m • i = 1 (the first natural frequency is close to 2 hz) • ξ1, damping ratio of the first mode shape = 0.0143 (locked tmd), 0.06 (free tmd) 4.1. locked tmd the peak accelerations at midspan amov using moving pedestrian load model according to eq. (19) was figure 3. response of the structure stationary pedestrian load, locked tmd. figure 4. response of the structure moving pedestrian load, locked tmd. evaluated and can be seen in figure 4. the value for reduction coefficient was calculated from eq. (18). calculation of r for locked tmd: r = −6.248 l + 0.0495i + 14.42ξ1 + 0.726 r = −6.248 17.4 + 0.0495 × 1 + 14.42 × 0.0143 + 0.726 r = 0.623 the amplitude of the stationary pulsating force is then r × 280 = 174 n. applying this force according to eq. (19) in the midspan of the beam, the maximal acceleration ast was evaluated. 118 vol. 15/2018 relation between stationary and moving pedestrian load models figure 5. static scheme of experimental footbridge [5]. tmd r acceleration moving model [ms−2] acceleration stationary model [ms−2] locked 0.623 1.100 1.073 free 1 0.264 0.261 table 1. maximal acceleration at midspan. 4.2. free tmd the evaluation of acceleration with free tmd was performed using the same method as with locked tmd. due to high structural damping (≈6%), the reduction factor r exceeded limit value 1 which was presented in eq. (20). therefore, the value of r is set to 1. comparison of the results can be seen in table 1. 5. conclusions the formula for the reduction coefficient r, which can be used for simulation of single pedestrian load, was proposed. this formula simplifies the pedestrian load from moving force model to stationary harmonic force model but keeps the accuracy of moving model. the formula is applicable on footbridges with static scheme of simply supported beams and with one of the first three natural frequencies close to 2 hz. the proposed method could be a very effective way to evaluate response of the structure to a single pedestrian load, because it does not require a time analysis and can be simply performed in commercial static software. parametric study in section 3 proved it is mainly length of the bridge and structural damping that influences the reduction coefficient. the numerical analysis in section 4 proved the estimation of reduction factor r = 0.6 0.7 [1] for low damped structures. however, it also showed that these approximations can lead to underestimation of load, which was demonstrated on the structure with tmd in this case. verification of the proposed formula was performed with very accurate results. nevertheless, further experimental verification and investigation in this field shall be performed. acknowledgements the authors gratefully acknowledge support from the czech technical university in prague, project sgs18/037/ohk1/1t/11 development and application of advanced algorithms for numerical analysis and modeling in mechanics of structures and materials references [1] štimac grandić. serviceability verification of pedestrian bridges under pedestrian loading. tehnički vjesnik 22:527–537, 2015. doi:10.17559/tv-20131030105641. [2] setra. footbridges: assessment of vibrational behaviour of footbridges under pedestrian loading, 2006. [3] fib bulletin 32: guidelines for the design of footbridges, 2005. [4] s. živanović, a. pavić, p. reynolds. vibration serviceability of footbridges under human-induced excitation: a literature review. journal of sound and vibration 279(1-2):1–74, 2015. doi:10.1016/j.jsv.2004.01.019. [5] a. dazio. fundamentals of structural dynamics. an-najah national university, 2013. [6] guide to basis of bridge design related to eurocodes supplemented by practical design handbook 4. leonardo da vinci pilot project, 2004. [7] e. caetano, a. cunha, w. hoorpah, j. raoul. footbridge vibration design. crc press, 2009. [8] e. shahabpoor, a. pavic, v. racic. interaction between walking humans and structures in vertical direction: a literature review. shock and vibration p. 22, 2016. doi:10.1155/2016/3430285. 119 http://dx.doi.org/10.17559/tv-20131030105641 http://dx.doi.org/10.1016/j.jsv.2004.01.019 http://dx.doi.org/10.1155/2016/3430285 acta polytechnica ctu proceedings 15:114–119, 2018 1 introduction 2 pedestrian load models 2.1 single pedestrian load 2.2 stationary pedestrian load model 2.3 moving pedestrian load model 3 parametric study 3.1 set of simply supported beams 3.2 response of the structure 3.2.1 response to stationary pulsating force 3.2.2 response to moving pulsating force 3.3 evaluation 4 verification of the proposed formula 4.1 locked tmd 4.2 free tmd 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0001 acta polytechnica ctu proceedings 12:1–4, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app an issue of intelligent road transport in kazakhstan kairatolla k. abishev∗, ruslan b. mukanov, asylbek zh. kasenov, almagul n. baltabekova s.toraighyrov pavlodar state university, pavlodar city, republic of kazakhstan ∗ corresponding author: kairatolla76@gmail.com abstract. the article reviewed the road transport sector of the republic of kazakhstan and found a significant role in ensuring road transport of freight and passengers. shown needs to improve the work of transport-communication complex of the republic of kazakhstan on the basis of using the latest information and communication technologies. described advantages of realization of several intellectual transport technologies. presented modern level of using satellite navigational systems on the automobiles, in addition written some automation and informational aspects of controlling transport processes. also presented measures to solve technological rearmament problem of the republic of kazakhstan. keywords: the republic of kazakhstan, automobile transport, intelligent transport systems, dispatching system. 1. introduction the republic of kazakhstan is entering a new phase in the reform of its economy, defining the purpose a significant increase in the efficiency of the national development system. message from the president of the republic of kazakhstan n.a. nazarbayev to the people has become a new strategic project on the path of accelerated development of the national economy, in close collaboration with the world of the global economy. presenting in his letter deeply thoughtful and detail-calculated economic and social strategy for the further successful movement of the people of kazakhstan to the prosperity, peace and the achievement of higher goals, the president gave her the name of "nurly zhol – path to the future". this program is intended to be the growth engine of the national economy in the upcoming years to improve the welfare and quality of life of every citizen in kazakhstan. according to the president, the most important area for the development of kazakhstan’s is the transport and logistics infrastructure. improving the transport sector is more than urgent for kazakhstan in conditions when there is a problem of intensification of trade and economic cooperation with europe and east asia. the introduction of intelligent transport systems will facilitate the organization of international transport, cargo tracking, and reduce the time of delivery, provide high performance and security of transport and logistics processes. 2. main part transport plays the most important role in the republic of kazakhstan life. on one hand, it is a very big territory (2725 thousand. km2), low density of its population (an average of about 6 people / km2), very big distances between sources of raw materials and production resources. in this situation, transport and communication complex provide not only the economic but also the political integrity of the country. on the other hand, the geographical location in the heart of the eurasian continent, far from the main labor and capital markets result in the inevitable growth of the transport component in the export-import operations.it consequently leads to a weak integration into the global economic system. the geopolitical location of kazakhstan between the capacious and dynamically developing markets of europe, east and south-east asia provides a chance to compensate this, by bringing transcontinental transit on the territory of the country[1]. 2.1. the road transport role in the common transport complex of the republic of kazakhstan as we can see the world and now the domestic experience, road transport is maximally adapted to function in today’s world. automobile transport is an inseparable and a very important component of transport significant role in transport and communication complex of the republic of kazakhstan. among its most important, the following characteristics can be outlined: a. popularity and wide availability, b. high maneuverability and speed of passenger transportation and goods delivery, c. possibility of door-to-door delivery without the necessity of any additional interim operations with goods, 1 http://dx.doi.org/10.14311/app.2017.12.0001 http://ojs.cvut.cz/ojs/index.php/app k.k.abishev, r.b.mukanov, a.zh.kasenov, a.n.baltabekova acta polytechnica ctu proceedings d. providing of near to zero-option short-distance service (delivery within the city and to the suburbs, and also those in rural areas), e. high degree of adaptation to different technological processes – both in industry and service sector, f. relatively low capital capacity. given the conditions mentioned above, the automobile transport is considered an inseparable component of all modern transport technologies including the integrated and multimodal transportation. the goods transportation activity in the republic of kazakhstan follow increasing trends in terms of capacity of delivered goods. according to rok statistics agency [2], 3 627,9 million tons of goods have been shipped via transport networks of the republic of kazakhstan from january to december in 2014, which is 3.7% more than within the same period in 2013. the cargo turnover for the given period was evaluated as equal to 487.4 billion t/km, 21293,00 billion pf passengers have been transported (which is 6.5% more than during january-december 2013), passenger turnover was 249,6 billion t/km (increase per 6.8% compared to 2013). according to the agency data, the proportion of automobile transport in relation to both transportation volumes and goods turnover has been increasing each year. currently, the automobile segment occupies 86,23% from total amount of delivered goods, 7,54% belong to railway transport and 6.09% – to pipelines, and 0.14% – to the other means of transport (air, internal water transport) – see figure1. figure 1. distribution of total transportation volumes among the means of transportation. as shown by international and domestic experience, automobile transport is ultimately adapted to be functioning in the conditions of market demand and supply. it can be proved by the fact that transport and communication complex was the primary choice for realization of privatization[3]. 2.2. necessary improvement for transport and communication complex of the republic of kazakhstan at this level of social development, there is a need for technological improvement of the transport process caused by a number of additional factors. first of all, the increased density of vehicular traffic as a result of development of economic and cultural ties and the accompanying increase in the process of trade flows and mobility. in these conditions, for this time the intensification of all areas of human activity is becoming increasingly important factor of time and therefore introduces the task to improve traffic speed. meanwhile, the transport infrastructure capabilities tend to lag behind the need to provide quality transportation in the face of increased intensity and speed. this often leads to the occurrence of accidents involving property damage, injuries and deaths. the traffic growth creates many problems for instance, traffic control, associated with the need to make responsible decisions in conditions of high traffic density of vehicles that often becomes one of the causes of accidents caused by human error. from this perspective, the technological improvement of the transport process is essential and urgent in order to achieve goals for the society improving the safety of vehicles[4]. thus, the introduction of modern information and communication technologies in transport significantly reduces the influence of the human factor[5][6]. 2.3. intelligent transport systems in the republic of kazakhstan intelligent transport systems – this is the system integration of modern information and communication technology and automation with transport infrastructure, vehicles and users, focused on safety improvment and efficiency of the transport process, comfort for operators, drivers and transport users. the introduction of intelligent technologies in the transport sector of kazakhstan made in the individual modes of transport and aimed at solving corporate problems. realization of certain intelligent transport technologies, and finally the formation of intelligent transport systems both have social and economic objectives, they are: a. traffic safety improvement and, as a consequence, the reduction of social wastes in the transport sector b. reduction in transport costs in the economy and other spheres of transport services, c. improvement of the environment and saving human and material (f.e. energy) resources. the strategic aim of the creation of intelligent system of road transport is the creation in the city: republics of interconnected intelligent systems that control traffic of transport vehicles and special emergency services, housing management. the strategic goal can also determine the formation of the republic competitive "road corridors", equipped with priority its technologies: movement controlling systems, traffic management systems that provide 2 vol. 12/2017 an issue of intelligent road transport in kazakhstan figure 2. ansdu work scheme, based on the satellite navigation. driver information about road surface, traffic conditions, the recommended routes of driving in the real situations on the roads, the availability and location road service areas, roadside services and other[5]. one of the main directions of development of transport technologies in the passenger transport sector of the republic of kazakhstan is the introduction of computer-based navigation systems of dispatching management. these systems use the location of vehicles with the signals of global navigation systems gps and glonass. dispatching system based on satellite navigation systems allow operational management of transportation, fixing transport work by the transmission and processing of information about the location of vehicles, access to that information for all interested participants of the transport process. schematic diagram of the operation of the automated dispatch control navigation systems based on satellite navigation is shown in figure2. installation of navigation systems on public passenger transport are focused not so much on the movement of the vehicle itself, but rather on comfort and safety of passengers[7]. 3. conclusion the need of the introduction of modern innovative technologies in all parts of the economy is one of the circumstances to improve its efficiency in the republic of kazakhstan. it is fully recognized by the government at the highest level and has repeatedly emphasized on public statements of the president of republic of kazakhstan. as for technological modernization of the transport industry of the republic of kazakhstan, the following tasks should be solved: (1.) formation of a single information space, (2.) training of qualified personnel to work in the new technological environment, (3.) the formation of regulatory and legal framework, standardization and harmonization of requirements for technical solutions in the field of information support of transport activities. solving these problems is very important to improve the operation of transport. references [1] m. bekmagambetov, r. zhumagulov. the transport system of kazakhstan in modern conditions. almaty, 2008. [2] http://www.stat.gov.kz, last accessed may 2016. [3] k. k. abishev, r. b. mukanov. improving road traffic safety in the republic of kazakhstan (rok). driver-car interaction and safety 2014, pg. 43-46. [4] m. novák, z. votruba, m. svítek, p. bouchner. improvement of bus and truck driving safety. (2007) conference proceedings – ieee international conference on systems, man and cybernetics, 1, art. no. 4273847, pp. 310-315. 3 http://www.stat.gov.kz k.k.abishev, r.b.mukanov, a.zh.kasenov, a.n.baltabekova acta polytechnica ctu proceedings [5] m. m. bekmagambetov. intelligent transport systems in the republic of kazakhstan. almaty, 2013. [6] p. bouchner, j. faber, m. novák, et al. human decision aspects in interactions with artificial systems. (2009) proceedings of the 2009 euro american conference on telematics and information systems: new opportunities to increase digital citizenship, eatis ’09, art. no. 1551749. [7] v. m. vlasov. transport telematics in the road sector. moscow-madi, 2013. 4 acta polytechnica ctu proceedings 12:1–4, 2017 1 introduction 2 main part 2.1 the road transport role in the common transport complex of the republic of kazakhstan 2.2 necessary improvement for transport and communication complex of the republic of kazakhstan 2.3 intelligent transport systems in the republic of kazakhstan 3 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0007 acta polytechnica ctu proceedings 7:7–11, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of virtual fields method, parallel network material model and finite element updating for material parameter determination florian dirisamer∗, umut d. çakmak, imre kállai, martín machado, zoltán major institute of polymer product engineering, johannes kepler university linz, altenbergstrasse 69, 4040 linz, austria ∗ corresponding author: florian.dirisamer@jku.at abstract. extracting material parameters from test specimens is very intensive in terms of cost and time, especially for viscoelastic material models, where the parameters are dependent of time (frequency), temperature and environmental conditions. therefore, three different methods for extracting these parameters were tested. firstly, digital image correlation combined with virtual fields method, secondly, a parallel network material model and thirdly, finite element updating. these three methods are shown and the results are compared in terms of accuracy and experimental effort. keywords: dic, finite element updating, parallel network model, vfm. 1. introduction elastomers are often used as damping elements and undergo, depending on their constraints, uniand multiaxial loadings in tension, compression and shearing [1–4]. the nonlinear inherent viscoelastic material behaviour of the epdm shore a 50 with a maximum elongation of several hundred percent makes it further complicated to extract material parameters for the whole range of time (frequency), temperature and environmental conditions relevant for the application [5, 6]. therefore a large number of tests at different temperatures and loading velocities have to be performed (75 tests [7] for the standard method (std)). to circumvent these difficulties and to extract the material parameters sufficiently, three different and well-known methods were used here for comparison purposes. the motivation is the biaxial characterization of elastomers for different applications (sealing, damping, etc.). firstly, a combined methodology based on digital image correlation (dic) [8] and virtual fields method (vfm) [9]. in this method, a heterogeneous strain field is introduced in order to acquire more than one parameter in a single test. this leads to drastically reduced time and costs. secondly, a parallel network material model (pnmm) with the procedure of bergström [10] (mcalibration and polyumod), where a test procedure including hysteretical and relaxation behaviour at certain temperatures, velocities, as well as test conditions leading to the modeling of the viscoelastic behaviour. thirdly, a finite element updating (feu) procedure using the abaqussolver [11] is utilized. measured force-displacement signals are compared to the simulated ones and the material parameters are updated until these two signals fit sufficiently within a defined tolerance. finally, these three methods are compared in terms of accuracy 36 80 120 r22 (a) . specimen 0 100 200 300 400 500 600 0 5 10 15 20 25 30 35 40 45 50 55 time in s d is pl ac em en t in m m 0.1 mm s−1 1 mm s−1 10 mm s−1 hyst. & res. (b) . displacement-time diagram. figure 1. biaxial test specimen with dimensions in mm and displacement-time diagram for 0.1 mm s−1, 1 mm s−1, 10 mm s−1 and with the procedure including hystereses and relaxation for the experiments. and experimental effort and the advantages as well as disadvantages are summarized for each method. 2. description of the methods the biaxial test specimens (figure 1a) were tested with a biaxial test frame [2] at 20 °c in a temperature chamber (cts gmbh, jennersdorf, aut), loaded by an 1.5 kn actuator and measured with a 4.5 kn load cell (bose electroforce systems group, mn, usa). the optical strain measurement was performed with a dic-camera system (aramis, gom mbh, braunschweig, d) and a cold light source (dedocool, dedotec usa inc., ma, usa) and a speckle pattern was manually applied to the biaxial test specimen. the further workflow for each of the three methods, dic combined with vfm (figure 4), pnmm with the procedure of bergström (mcalibration and polyumod) (figure 7) and feu (figure 10), is explained in detail in the subsections 2.12.3. 7 http://dx.doi.org/10.14311/app.2017.7.0007 http://ojs.cvut.cz/ojs/index.php/app f. dirisamer, u.d. çakmak, i. kállai et al. acta polytechnica ctu proceedings (a) (b) (c) (d) (e) (f) εxx εyy εxy εm ajor εm inor εres = εm inor εm ajor legend εxx legend εyy legend εxy legend εm ajor legend εm inor legend εres -0.1 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 -0.1 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 -0.2500 -0.1875 -0.1250 -0.0625 0.0000 0.0625 0.1250 0.1875 0.2500 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.000 0.025 0.050 0.075 0.100 0.125 0.150 0.5 0.6 0.7 0.8 0.9 1.0 figure 2. digital image correlation results at max strains in εxx (a), εyy (b) and εxy (c) and calculated results for εm ajor (d), εm inor (e) and εm inor/εm ajor (f). 0 10 20 30 40 50 0 1 2 3 strain in % st re ss in m p a 0.1 mm s−1 1 mm s−1 10 mm s−1 figure 3. stress-strain diagram for 0.1 mm s−1, 1 mm s−1 and 10 mm s−1. 2.1. digital image correlation (dic) and virtual fields method (vfm) the workflow for the vfm is shown in figure 4. displacement controlled tests at velocities of 0.1 mm s−1, 1 mm s−1 and 10 mm s−1 (figure 1b) were performed with optical strain measurement and the loaddisplacement data for each of the three velocities, as well as the optical images were recorded. these image series are evaluated via dic (ncorr, georgia institute of technology, ga, usa) to generate the strain fields εxx, εyy and the shear field εxy (figure 2a-c). afterwards εm ajor (figure 2d) and εm inor (figure 2e) were calculated according to equation 1. εm ajor/m inor = εxx + εyy 2 ± √( εxx + εyy 2 )2 − ( εxx · εyy − ε2xy ) (1) the equibiaxial area is the region, where εm ajor and εm inor are “equal”, therefore, εres = εm ajor/εm inor (figure 2f) was calculated and the equibiaxial area was extracted in the range of εres ≥ 0.5. that area perform test capture images image series matlab (dic) strain field matlab major/minor strain matlab stress-strain data load-displacement data figure 4. workflow for the virtual fields method (vfm). was afterwards converted to an equibiaxial diameter and the biaxiality ratio φ, the biaxial stress σ was calculated by dividing the measured force with the equibiaxial diameter and the thickness and the strain ε was calculated for each image series (figure 3). 8 vol. 7/2017 comparison of vfm, pnmm and feu 0 10 20 30 40 50 60 0 1 2 3 strain in % st re ss in m p a 0 20 40 601 . 4 5 1 . 7 5 time in s s tr es s in m p a 0 20 40 601 . 9 5 2 . 3 time in s s tr es s in m p a 0 20 40 602 . 2 2 . 5 5 time in s s tr es s in m p a 0 20 40 602 . 3 5 2 . 7 5 time in s s tr es s in m p a figure 5. stress-strain diagram for load with hystereses and relaxation with marked relaxation areas. parameter value parameter value bulk modulus 106.59 lock stretch 4.0841 pressure dependence of flow −0.129 temp. exp. flow resistance 0 temp. factor for stiffness 0 therm. expansion coefficient 0 thermal expansion ref. temp. 293 flow resistance a 0.1743 shear modulus a 7.9898 stress exponential a 11.499 final shear modulus b 0.9803 flow resistance b 1.4018 initial shear modulus b 2.0318 norm. evolution rate of µb 10.212 stress exponential b 10.156 relative contribution of i2 c 9.3e-05 shear modulus c 1.3068 table 1. parameter table for three network material model. 2.2. parallel network material model (pnmm) the workflow for the pnmm is shown in figure 7. for the procedure of bergström (mcalibration and polyumod), tests which include hysteresis and relaxation domains (figure 1b) was performed. the hysteresis and relaxation domains are necessary to model the viscoelastic behaviour with one material test at a certain temperature and velocity. as a material model, the three parallel network model [12] was chosen. it consists of: 1st network: temperature-dependent 8-chain model [13, 14], power-law flow rate, hydrostatic pressure dependence, power-law temperature dependence strain in % 0 10 20 30 40 500.0 0.5 1.0 1.5 2.0 2.5 3.0 st re ss in m p a experimental predicted figure 6. experimental and predicted stress-strain diagram. perform test load-displacement data abaqus stress-strain data mcalibration & polyumod material model material model parameter figure 7. workflow for the parallel network material model (pnmm). 2nd network: temperature-dependent 8-chain model, shear modulus evolution with plastic strain, power-law flow rate, hydrostatic pressure dependence, power-law temperature dependence 3rd network: 8-chain model with first order i2 (second invariant of deformation tensor) dependence. with a simulation, the stress-strain data (figure 5) was calculated, which is imported and processed in mcalibration. the material parameters are calculated analytically by comparing the predicted results to the experimental ones and updating these parameters until the results fit. in that way, the material parameters are generated (table 1) and the experimental as well as predicted strain-stress curves are shown in figure 6. 9 f. dirisamer, u.d. çakmak, i. kállai et al. acta polytechnica ctu proceedings 0 10 20 30 40 50 0 50 100 150 200 250 displacement in mm lo ad in n 0.1 mm s−1 1 mm s−1 10 mm s−1 experimental predicted figure 8. experimental and predicted forcedisplacement diagram with the mooney-rivlin model. velocity c10 c01 d[mm s−1] [mpa] [mpa] 0.1 0.5849 0.0856 0 1 0.5924 0.1016 0 10 0.5979 0.1506 0 table 2. parameter table of the received mooneyrivlin model-parameters. 2.3. finite element updating (feu) the workflow for the feu is shown in figure 10. tests with velocities of 0.1 mm s−1, 1 mm s−1 and 10 mm s−1 (figure 1b) were performed and the load-displacement data were recorded (figure 8, solid lines). these data were imported separately to matlab [15] and a custom-made feu-script is executed. this prepares the input file for abaqus automatically, executes the solver, performs a numerical simulation and extracts the simulated results. it compares the simulated (predicted) results to the experimental ones, updates the material parameters in the input file and repeats that procedure until the load-displacement curve fit within a defined tolerance. the gained material parameters are liste in table 2 and the compared load-displacement data are shown in figure 8. furthermore, the evolution of the parameters c10 and c01 are visualized in figure 9. 3. results and discussion in order to analyze the difference of the three terms of accuracy and experimental effort, the advantages as well as the disadvantages are summarized individually. 10−1 100 101 0.58 0.59 0.6 velocity in mm s−1 c 10 in m p a c10 0.06 0.11 0.16 c 01 in m p a c10 c01 figure 9. evolution of c10 and c01 over the velocity. perform test load-displacement data matlab abaqus material model parameter figure 10. workflow for the finite element updating (feu). • virtual fields method (vfm) the stress-strain data were calculated by vfm, in which the actual full-field strain data is considered. this full-field data includes some information about the sliding of the clamping system, elastic extension of the machine due to the force and some other small effects, which cause deviation between the measured and applied length. therefore, these effects can be taken into account and excluded. furthermore, more than one material parameter can be analyzed by generating inhomogeneous strain fields (different loading states) in one test and specimen. however, there are difficulties in finding the right equations for the virtual strain field of each loading condition and its specific (geometric) constraints to receive the right parameters and dic data is needed, which cause higher effort for acquisition and a huge amount of computational power. • parallel network material model (pnmm) the result of pnmm are the nonlinear material model parameter which include here (but is not limited to) viscoelastic, relaxation and cyclic behaviour, which gives the whole range of velocity dependence at a certain temperature and loading configuration with one test. stress-strain data must be evaluated by a simulation and the analytical calculation of the material model parameters is time consuming. furthermore, different parameter combinations of the same material model and the same test data can be received by choosing different initial parameters. therefore, the results have to be examined carefully. the selection of the material model is important for the characteristics and the accuracy. 10 vol. 7/2017 comparison of vfm, pnmm and feu std vfm pnmm feu 0 0.5 1 1.5 2 2.5 1 0.33 0.2 11 2 1.5 0.5 e ffo rt manual experimental computational figure 11. manual experimental and computational effort of virtual fields method (vmf), parallel network material model (pnmm) and finite element updating (feu) normalized to the standard method (std). • finite element updating (feu) the result of feu are the nonlinear material model parameters. therefore, results from specimen, as well as component tests can be taken to create a simulation. to receive the appropriate results, it is necessary to create a detailed simulation with the proper boundary conditions. furthermore, the selection of the material model is essential for the results. however, there are different parameter combinations for the same material model and simulation just by choosing different initial parameters. careful analyses of the results at various loading rates are required. the comparison of the manual and computational experimental effort of the three methods vfm, pnmm and feu, normalized to the standard method (std) (cf. section 1) is shown in figure 11. the manual experimental effort for the standard test consists of three test procedures (uniaxial, pure shear, equibiaxial), five temperatures and five test velocities, which leads to 75 tests (section 1). through the inhomogeneous strainfield of the vfm, the three test procedures can be minimized to one, which gives 25 tests, however, because of dic, the computational effort is more than doubled compared to std. for the pnmm, the whole range of velocity data for a certain temperature is included in one test, this cause 15 tests altogether, but the computational effort is by 50% greater than std. this is caused by the desired stress-strain data and the analytical calculations. the feu has no benefit on the reduction of the number of tests, nevertheless, the computational effort for small fem-simulations can be minimized, as the force-displacement data is selected directly. acknowledgements the research for this paper was performed within the framework of ffg 7th call of “produktion der zukunft” (project number: 848620). references [1] l. treloar. the elasticity and related properties of rubbers. reports on progress in physics 36(7):755, 1973. [2] u. d. çakmak, z. major. experimental thermomechanical analysis of elastomers under uniand biaxial tensile stress state. experimental mechanics 54(4):653–663, 2014. [3] u. d. çakmak, i. kallaí, z. major. temperature dependent bulge test for elastomers. mechanics research communications 60:27–32, 2014. [4] e. leiss-holzinger, u. d. çakmak, b. heise, et al. evaluation of structural change and local strain distribution in polymers comparatively imaged by ffsa and oct techniques. express polym lett 6:249–256, 2012. [5] n. w. tschoegl, w. g. knauss, i. emri. the effect of temperature and pressure on the mechanical properties of thermoand/or piezorheologically simple polymeric materials in thermodynamic equilibrium–a critical review. mechanics of time-dependent materials 6(1):53–99, 2002. [6] u. d. çakmak. experimental studies on time-temperature-dependent bulk and surface behavior of soft composite materials. ph.d. thesis, johannes kepler university, 2014. [7] a. grambow. bestimmung der materialparameter gefüllter elastomere in abhängigkeit von zeit, temperatur und beanspruchungszustand. ph.d. thesis, bibliothek der rwth aachen, 2002. [8] ncorr. ncorr. http://www.ncorr.com, 2016. [9] f. pierron, m. grédiac. the virtual fields method: extracting constitutive mechanical parameters from full-field deformation measurements. springer science & business media, 2012. [10] j. bergström. polyumod user’s manual. veryst engineering, llc, needham, ma 2009. [11] hibbett, karlsson, sorensen. abaqus/standard: user’s manual, vol. 1. hibbitt, karlsson & sorensen, 1998. [12] j. bergström, j. bischoff. an advanced thermomechanical constitutive model for uhmwpe. the international journal of structural changes in solids 2(1):31–39, 2010. [13] e. m. arruda, m. c. boyce. a three-dimensional constitutive model for the large stretch behavior of rubber elastic materials. journal of the mechanics and physics of solids 41(2):389–412, 1993. [14] j. bergström, m. boyce. large strain time-dependent behavior of filled elastomers. mechanics of materials 32(11):627–644, 2000. [15] c. moler, et al. matlab user’s guide. university of new mexico, 1982. 11 http://www.ncorr.com acta polytechnica ctu proceedings 7:7–11, 2017 1 introduction 2 description of the methods 2.1 digital image correlation (dic) and virtual fields method (vfm) 2.2 parallel network material model (pnmm) 2.3 finite element updating (feu) 3 results and discussion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.21.0005 acta polytechnica ctu proceedings 21:5–9, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app use of finely ground recycled concrete for improvement of interfacial adhesion in fiber-reinforced cementitious composites radim hlůžek∗, jan trejbal czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague, czech republic ∗ corresponding author: radim.hluzek@fsv.cvut.cz abstract. this paper deals with an improvement and an assessment of a polymeric macro-fibers adhesion to the cement matrix. for this purpose, two approaches were employed – (i) roughening of fibers using a plasma treatment and (ii) an addition of finely ground recycled concrete (amount 30 wt. %) to the matrix ensuring the roughness of interfacial zones. polyethylene terephthalate (pet) and polypropylene (pp) fibers, both ca. 0.3 mm in a diameter, were used. these were surface roughened using a cold oxygen plasma treatment and then observed by means of the scanning electron microscopy. consequently, pull-out tests of an individual fiber embedded 25 mm in the matrix were performed, while the force needed for fiber pullout was recorded. results have shown that plasma treated fibers reached on a better adhesion with the matrix by up to ca. 5 % (pet) and 20 % (pp), if compared to reference fibers. when recycled concrete was used, the adhesion increased further by about 5–10 % for both fiber types. keywords: recycled concrete, waste, adhesion, fiber reinforcement, fibrous composite. 1. introduction it is widely known that waste production shows an unsustainable trend in global scale. this also hold true for the construction industry. it is reported that global production of construction demolition waste ranges from 3 to 10 billion tonnes per year of which 3.5 million tonnes in the czech republic. these basic data provide the sufficient proof of why it is necessary to deal with the reduction of waste and its re-use. therefore, a law on a total ban on landfilling of inert waste with effect from 2024 was ready within the european union [1]. it is therefore necessary to find new efficient solutions for the reprocessing of construction waste including concrete from demolished buildings [2]. in civil engineering praxis, waste concrete is crushed and aggregate is then separated and re-used within production of new concrete mixtures. but the question of the use of dust emerging during crushing and grinding processes arises [3]. such dust may contain up to 10 % of un-hydrated cement grains [4]. these can be exposed and isolated from dust using a highspeed milling [5]. it is possible to use such obtained material as replacement of cement at the form of socalled active micro aggregate. it is appropriate for production of lightweight concrete blocks, mortars, etc. [4–6]. however, the potential of this material is still limited due to particles different thickness and size and presence of various impurities [4, 7, 8]. some researches tried to use finely ground recycled concrete for filling of pores in various cement-based composites. individual particles can fill also spaces in interfacial zones in fibrous composite materials and thus contribute to the better adhesion between reinforcing fibers and the matrix [9, 10]. unfortunately, such phenomenon has not been so far clearly explained, tested and described. therefore, we tested the adhesion between smooth and roughened fibers and the cement matrix containing finely ground recycled concrete by means of pull out tests of individual fibers from specimens made from both matrices. 2. materials 2.1. fibers two types of polymeric macro-fibers were used: (i) polyethylene terephthalate (pet) and (ii) polypropylene (pp). these were made by the czech manufacturer spokar (plc). both types were surface treated using cold oxygen plasma to roughen and to activate their surface. this process has been described in more details in our previous study [11]. fibers are further marked as follows: pp and pet, reference; pp-t and pet-t, treated. the basic geometric and mechanical properties of fibers are summarized in table 1. 2.2. cement and finely ground recycled concrete two different types of matrices were made. the first one was composed only from portland cement (cem i 42.5 r) with water to cement ratio (w/c) equal to 0.4 (marked cem), while the second one contained finely ground recycled concrete in amount of 30 % of weight of cement (marked rec), water to binder 5 http://dx.doi.org/10.14311/app.2019.21.0005 http://ojs.cvut.cz/ojs/index.php/app radim hlůžek, jan trejbal acta polytechnica ctu proceedings fiber type diameter density young’s modulus tensile strength [µm] [kg/m3] of elasticity [gpa] [mpa] pet / pet-t 335 950 5.8 238 pp / pp-t 305 900 6.1 440 table 1. basic geometric and mechanical properties of fibers [12]. (w/b = w/(cem+rec)) was increased to 0.41 (increased due to the high specific surface area of recycled concrete). waste recycled concrete used in this study, originated from railway sleepers, was ground using high-speed milling technology with maximum grain size of 0.25 mm. the hardening and curing of all test produced specimens took place in standard laboratory conditions for 28 days. 3. experimental methods 3.1. scanning electron microscopy morphology of fibers was observed using the merlin zeiss scanning electron microscopy device. in order to perform this analysis, it was necessary to provide sufficient electrical conductivity of the fiber surface. therefore, all tested fibers were coated by a thin layer of gold using the boc edward scancoats six machine. process parameters were: 1.3 kv voltage, 35 ma electric current, and 40 s plating time. it was assumed that such a treatment had not any effect on physical properties of fiber surfaces. 3.2. pull-out tests prismatic samples having dimensions of 25 × 20 × 25 mm were made from both types of matrices. the single fiber was always placed in the center of the specimen with a embedded length equal to the prism height, i.e. 25 mm. an illustrative specimen is shown in figure 1. pull-out tests were performed 28 days after hardening and curing of mixtures. in the first place, each tested sample was clamped into the static jaw of the hydraulic loading frame (web tiv ravestein fp100). secondly, the free end of the fiber was clamped into the opposite sliding jaw of the press. finally, the fiber was gradually pulled out from the sample at a constant rate of 2 mm/min. during the process, the free-end displacement and resisting force were measured. the loading process was terminated when the displacement reached 5 mm. such obtained data were evaluated using the dipro software based on the matlab software. the adhesion between the two materials was then expressed by interfacial shear stress (1) and (2), which is commonly used in the field of fiber reinforced composites because it provides values normalized to the fiber surface area. the stress was calculated in two stages: after reaching a fiber free-end pull out of 3.5 mm (hereinafter referred to as “3.5”) and when the measured force achieved on its maximum (denoted by the “max” index). figure 1. prepared specimen for the pull-out test. τmax = fmax cf le (1) τ3.5 = f3.5 cf le (2) f3.5 and fmax represent the force necessary to reach the 3.5 mm fibers free-end displacement and the maximal reached force, respectively. cf is a fiber circumference and le is the embedded fiber length. 4. results 4.1. scanning electron microscopy a scanning electron microscopy image of the reference pet fiber (figure 2) revealed its smooth surface, which is very inappropriate from the mechanical point of view. compared to that, the same fiber after 120 seconds of the plasma treatment was already gently roughened by irregular holes, as shown in figure 2. this desirable morphology change was caused by surface etching and the ion bombardment. it should be noted that the roughening of the treated fibers surface was observed only after 120 seconds. if the fibers were exposed to plasma for shorten time, there were no visible changes on their surface. similarly to pet fibers, the surface of reference pp fibers was too smooth, except for longitudinal profiling apparently originates from production technology (stretching process), as shown in figure 3. after 120 seconds lasting modification, minor changes – longitudinal recesses along the length of the fiber – in morphology were observed, see figure 3. it can be 6 vol. 21/2019 recycled concrete in fiber-reinforced cementitious composites assumed that these tiny morphology changes will not affect the adhesion of fibers to the matrix. extending the exposure time would certainly increase the effect of roughening, but cross-section of fibers significantly decreases. 4.2. pull-out tests based on pull-out test results, it can be observed that the highest value of the interfacial shear stress was recorded in the case of plasma modified pp fibers in combination with the matrix containing finely ground recycled concrete (rec). as shown in figure 5, the interfacial shear stress at maximum force was about 0.45 mpa and 0.38 mpa at the 3.5 mm fibers free-end displacement, respectively. in the case of cementitious matrix with same fibers, results were as follows: ca. 0.4 resp. 0.34 mpa. furthermore, it should be noted that for untreated fibers, values of the interfacial shear stress with respect to the matrix used did not differ substantially. in the case of pet fibers, the same trend as for pp fibers mentioned above can be observed, but only with significantly smaller differences, see figure 4. unlike the assumption, it has been found that the adhesion between reference pet fibers in combination with the matrix containing recyclate (rec) was a little bit worse than between same fibers and the matrix containing only cement (cem). the maximum interfacial shear stress dropped from ca. 0.31 mpa (cem) to 0.28 mpa (rec) and at the 3.5 mm displacement from 0.31 to 0.3 mpa. nevertheless, in the case of the use of plasma treated fibers, the adhesion was increased, especially in combination with the matrix containing recycled concrete. the maximum shear stress was increased by ca. 5–10 % up to ca. 0.33 mpa. to sum up, it can be stated that the use of finely ground recycled concrete as a partial cement additive can improve the adhesion between polymeric fibers and the cement-based matrix, but only in the case of the use of plasma treated fibers. in the case of reference fibers with a smooth surface, the use of recycled material appears to be ineffective. thus, physical interaction between fine particles of the recycled material and the plasma-roughened fiber surface plays a significant role in achieving the desired increase in the adhesion between the fiber and the matrix. 5. conclusion this study researched the adhesion between polymeric macro fibers (polyethylene terephthalate and polypropylene), having ca. 0.3 mm in the diameter, and two types of cement matrices. the reference matrix was composed of cement (cem i 42.5 r, w/c 0.4), while in the second case, 30 wt. % of cement was replaced with recycled concrete at the form of finely ground powder. in order to roughen fiber surfaces and thus to ensure the adhesion between them and the matrix, fibers were treated by cold oxygen plasma. changes on fiber surfaces were observed using the sem microscopy. interaction between the two materials was assessed directly by means of the single fiber pull-out test from matrix specimens (25 × 20 × 25 mm). it was shown that the surface roughening of pp fibers was much more intensive than in case of pet fibers. pull-out tests revealed that the adhesion between pp and pet fibers and the cement matrix was constant regardless to the fiber treatment. similarly, when these fibers were pulled out from the matrix containing finely ground concrete powder, the adhesion was not increased, if compared to reference materials. on the other hand, treated fibers (both pp and pet) showed significantly increased adhesion with the matrix containing recycled concrete. this phenomenon was attributed to fiber roughening and to working of recycled concrete grains in interfacial zones. acknowledgements this work was financially supported by the czech technical university in prague – sgs project sgs18/037/ohk1/1t/11 and the czech science foundation [grant number ga čr 17-06771s]. the authors would also like to thank to tereza horová for her help with this research. references [1] m. škopán. analysis of the current state of c&dw recycling in the czech republic and assessment of development options. recycling 129(9):19–27, 2005. [2] m. d. schepper, r. snellings, k. d. buysser, et al. the hydration of cement regenerated from completely recyclable concrete. construction and building materials 60:33–41, 2014. doi:10.1016/j.conbuildmat.2014.02.067. [3] t. pavlů, m. šefflová. environmental assessement of the structural elements with recycled aggregate. recycling 129(5):108–112, 2005. [4] j. xiao, j. li, c. zhang. mechanical properties of recycled aggregate concrete under uniaxial loading. cement and concrete research 35(6):1187–1194, 2005. doi:10.1016/j.cemconres.2004.09.020. [5] l. evangelista, j. de brito. mechanical behaviour of concrete made with fine recycled concrete aggregates. cement and concrete composites 29(5):397–401, 2007. doi:10.1016/j.cemconcomp.2006.12.004. [6] j. topič, j. trejbal, t. plachý, z. prošek. relationship between compressive strength and young’s modulus of cement paste with recycled concrete powder. key eng mater 722 kem:254–259, 2017. doi:10.4028/www.scientific.net/kem.722.254. [7] v. růžička. lightweight blocks based on recycled concrete. ph.d. thesis, czech technical university in prague, faculty of civil engineering, 2017. [8] m. etxeberria, e. vázquez, a. marí, m. barra. influence of amount of recycled coarse aggregates and production process on properties of recycled aggregate concrete. cement and concrete research 37(5):735–742, 2007. doi:10.1016/j.cemconres.2007.02.002. 7 http://dx.doi.org/10.1016/j.conbuildmat.2014.02.067 http://dx.doi.org/10.1016/j.cemconres.2004.09.020 http://dx.doi.org/10.1016/j.cemconcomp.2006.12.004 http://dx.doi.org/10.4028/www.scientific.net/kem.722.254 http://dx.doi.org/10.1016/j.cemconres.2007.02.002 radim hlůžek, jan trejbal acta polytechnica ctu proceedings figure 2. sem images of reference (left) and modified (right) pet fiber. figure 3. sem images of reference (left) and modified (right) pp fiber. figure 4. interfacial shear stress values for different combinations of matrices and pet fibers calculated from the pull-out test. 8 vol. 21/2019 recycled concrete in fiber-reinforced cementitious composites figure 5. interfacial shear stress values for different combinations of matrices and pp fibers calculated from the pull-out test. [9] n. sebaibi, m. benzerzour, n. e. abriak, c. binetruy. mechanical properties of concrete-reinforced fibres and powders with crushed thermoset composites: the influence of fibre/matrix interaction. construction and building materials 29:332–338, 2012. doi:10.1016/j.conbuildmat.2011.10.026. [10] k. r. akça, ö. çakır, m. i̇pek. properties of polypropylene fiber reinforced concrete using recycled aggregates. construction and building materials 98:620– 630, 2015. doi:10.1016/j.conbuildmat.2015.08.133. [11] r. hlůžek, z. prošek, j. trejbal, et al. plasma treatment impact on physical and chemical properties of polymeric fibers. acta polytechnica ctu proceedings 13:49–54, 2017. doi:10.14311/app.2017.13.0049. [12] r. hlůžek. analysis and modification of interfacial zones between cementitious matrix and fiber reinforcement. ph.d. thesis, czech technical university in prague, faculty of civil engineering, 2018. 9 http://dx.doi.org/10.1016/j.conbuildmat.2011.10.026 http://dx.doi.org/10.1016/j.conbuildmat.2015.08.133 http://dx.doi.org/10.14311/app.2017.13.0049 acta polytechnica ctu proceedings 21:5–9, 2019 1 introduction 2 materials 2.1 fibers 2.2 cement and finely ground recycled concrete 3 experimental methods 3.1 scanning electron microscopy 3.2 pull-out tests 4 results 4.1 scanning electron microscopy 4.2 pull-out tests 5 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.23.0038 acta polytechnica ctu proceedings 23:38–43, 2019 © czech technical university in prague, 2019 available online at https://ojs.cvut.cz/ojs/index.php/app evaluation of underground structures subjected to seismic loads jan pruška czech technical university in prague, faculty of civil engineering, department of geotechnics, thákurova 7, 166 29 prague, czech republic correspondence: pruska@fsv.cvut.cz abstract. the paper is focused on the evaluation of the effect of earthquakes on underground structures. free-field analysis is one solution of this task common mainly in engineering tunnelling practice, but it has some rather simplified aspects (e.g. equivalent shear strain is constant). pseudostatic finite element calculation combines free-field analysis and the advantages of a fem model. dynamic effects are introduced in the form of displacements prescribed along the vertical boundaries of the fem model in a usually static manner. this approach also implies constant material parameters for the geological profile in the horizontal direction, an arbitrary geometry of excavation, soil structure interaction and description of share waves as a time-dependent 1d analysis of the so called free-field column. moreover, there is shown an example comparing pseudo-static fem analysis with an analytical method. finally, the advantages of the pseudo-static fem method are presented. keywords: finite element method, earthquake, 1d free field analysis, dynamics, pseudo static analyses. 1. introduction the response of underground structures subject to seismic load can be solved in several ways, which can be roughly classified into four categories: laboratory testing, analytical solutions, pseudo static finite element calculations and full dynamic analysis. (1.) centrifuge modelling techniques allow well for control testing conditions regarding investigations of soil structure interaction in simulated seismic events. however, there are several difficulties for a dynamic centrifuge model such as accurate modelling of the boundary conditions or the scaling law for dynamic velocity. (2.) most simplified methods were derived from closed form solutions but involving neglect of some factors. in general, the methodology of the analytical method is that the seismic loads for an underground structure are characterized in terms of strains imposed on the structure by the surrounding ground or their interaction. although this is the most conservative method having only a limited application, it does provide a first-order estimate of the liming stresses so to potentially serve as an accuracy check on more complex calculations. (3.) a pseudo-static finite element calculation provides a solution for any shape of underground structure in a generally non-homogeneous layered rock/soil mass with a potentially nonlinear material response. the basic idea of this approach is that the maximum shear deformation in a free-field condition of the layer represents the maximum dynamic earthquake stress in this layer and can be an input to the numerical model as a boundary condition. the shear strain equivalent to the actual dynamic loading can be calculated from the particle velocity estimated either in a simplified manner based on limited information or from the solution of a one dimensional simulation of free-field response to an actual earthquake. (4.) full dynamic analyses are based on a direct application of recorded accelerations to the actual computational model of the tunnel where the soil and tunnel responses are mechanically coupled and analysed via numerical modelling, such as finite element methods. 2. analytical method the methodology of the analytical method for a seismic loading design is based on the principle that the static condition has to include the additional loading imposed by seismicity. in the case of underground structures, the seismic loads are described in terms of strains imposed on the structure by the surrounding ground mass or their interaction. 2.1. seismic coefficient method this method was originally developed for dynamic analysis of over ground buildings. inertial forces due to earthquakes f1 of the structure and f2 of the upper soil mass are described by the following equations: f1 = a g · q = kc · q, (1) f2 = ηc · kh · m1 · g. (2) this method is not suitable for tunnels with large diameters because this method does not reflect the real dynamic property of tunnel structure. 38 https://doi.org/10.14311/app.2019.23.0038 https://ojs.cvut.cz/ojs/index.php/app vol. 23/2019 underground structures subjected to seismic loads 2.2. free – field deformation approach tunnel structures are assumed to move together with the surrounding soil/rock mass during earthquakes and therefore undergo the following deformations: • ovaling, • axial, • curvature. the ovaling deformation of a circular tunnel lining is primarily caused by seismic waves that propagate perpendicular to the longitudinal axis of the tunnel (figure 1). the vertically propagating shear waves produce usually the most critical ovaling distortion of the lining. the free field deformation approach assumes that the deformation of the underground structure should conform to the deformation of the soil in the free field under seismic shaking. in these simplified methods described in [1–4] the soil – structure interaction and the effects of the waves are neglected. moreover the maximum shear strain γmax (equation 6) induced by shear waves propogating in vertical planes is constant along the depth z. this may lead to an overestimate of tunnel lining deformations especially in loose rock and soils. figure 1. ovaling deformation of a circular cross section. in the case of a circular tunnel in the homogeneous medium the ovaling deformation is determined by the following steps: (1.) according to the seismic report we get the pick ground acceleration (3) and the site-specific pickground-acceleration (4): agr = 40% · ss, (3) amax,s = s · agr (4) (2.) and then we determine the pick acceleration at the depth of the tunnel: az,max = c · s · agr, (5) (3.) az,max max is used to determine maximum shear deformation in free-field condition as shown in equations (6) and (7): γmax = vz cs , (6) vz = k · az,max (7) (4.) and finally we obtained the ovaling deformation of the circular tunnel: ∆dff = γmean d 2 . (8) 2.3. soil-structure interaction approach the kinematic interaction between the tunnel and the surrounding soil/rock mass is used to improve the accuracy of the widely used free-field methods. generally, analytical solutions for estimating soil-structure interaction for tunnels are based on the following assumptions: • the tunnel is circular; • the soil is an infinite, elastic, homogeneous and isotropic medium; • the tunnel lining is generally an elastic tube under plane strain conditions or it is based on the theory of an elastic beam on an elastic foundation; • the interface between the soil and the tunnel lining has a full-slip or a no-slip condition; • the soil-structure interaction effects operate in a quasi-static manner, ignoring any inertial interaction effect. there exist many approaches in the technical literature describing seismic design and analysis of underground structures using a soil structure interaction approach. for example: billota et al. [5] discussed four different closed-form elastic solutions to evaluate the maximum shear strain γmax (all approaches are based on the equilibrium of the soil column from the surface to the given depth z), while expressions for displacements and internal lining forces in the circular tunnels excavated in a rock mass with a near fault are presented in [6], and closed-form expressions to calculate the circular tunnel liner forces due to the compressive seismic p-wave propagation are described by kouretzis et al. [7]. in addition an analytical solution for estimating internal forces in the lining of a circular tunnel in the homogeneous medium is published in [8, 9]. the ovaling deformation is determined by the assumed shear strain ∆d (figure 1). we get for parameters ∆dff, ∆dlin the follow-on formulas: ∆dlin = rn∆dff, ∆dff = γmean d 2 , rn = 4(1 − νm) αn + 1 ,αn = 12el(5 − 6) d3gm(1 − ν2l ) . (9) it is evident that the material and geometric parameters rn and αn take into account the rigidity of the subsoil-lining system. expressing the curvature of an elliptical excavation using ∆dlin and parameters describing soil structure interaction then the moment in the lining (figure 2) can be written using the polar 39 jan pruška acta polytechnica ctu proceedings coordinate ϕ (positive clockwise from horizontal) as: m(ϕ) = 6el∆dlin d2(1 − ν2l ) cos (2(ϕ + φ 4 )). (10) with reference to figure 2 (c), the normal force on the curved element can be expressed as: n(ϕ) = − 1 r dm2 dϕ + pnr (11) and after rearranging using equation (10): n(ϕ) = 48eli∆dlin d3(1 − ν2l ) cos (2(ϕ + φ 4 )) + rpn (12) figure 2. internal forces and contact stress on the differential element. 3. pseudo static fem implementation of free-field analysis results into the static fem analyses can be made in several ways. generally, the shear strains in the 2d solution without lining have to be in conformity with 1d free-field analyses and the shear deformation in the free-field deformation is applied by prescribed displacements at the boundaries of the numerical model. in the first method, we consider a homogeneous soil/rock mass with an effective shear modulus g. we determine the maximum shear deformation in free-field condition γmax (6). the value of γmax corresponds to the maximum horizontal displacement applied in the boundaries of the numerical model (figure 3), calculated by equation: ∆xmax = γmax · hm. (13) the second method considers the layered soil/rock mass (figure 4) where some sort of homogenization is needed. adopting the voigt assumption of constant deformation at each point of the solid we obtained effective shear modulus ghomv : ghomv = h1 h g1 + h2 h g2 (14) and average shear deformation of the numerical model γmean: γmean = h1 h γ1 + h2 h γ2. (15) figure 3. homogenized mass and constant shear strain. figure 4. layered mass and constant shear strain . for the layered mass the better method is based on the setup constant corresponding movements by parts – each layer has a different shear deformation – figure 5. the prescribed shear deformation values should be in the ratio of the stiffness of the materials in layers to ensure the continuity of the stress at the layer interface: γ1 γ2 ≈ g2 g1 (16) figure 5. layered mass with different values γ at layers 1 and 2. 4. full dynamic calculations using fem the numerical solution of dynamic load effects using the finite element method is based on the principle of virtual displacements, which express the external volume load by the effects of inertial forces and leads to equations of differential motion in the form: mü(t) + cu̇(t) + ku(t) = f (t). (17) 40 vol. 23/2019 underground structures subjected to seismic loads the assembly of the stiffness matrix c is usually a very complicated task thus the rayleigh formulation is used i.e. as a linear combination of the mass tensor m and the stiffness tensor k: c = αm + βk. (18) coefficients α and β have been calculated according to the double frequency method that avoids an overestimate of damping throughout the considered frequency range [10]. assuming that the first natural frequency is the least damping, the equation (17) can be decomposed to the natural vibrations. using the damping ratio ξ we get a relationship for individual frequencies (for real systems with a large number of degrees of freedom): ξi = 1 2 ( α ωi + βωi). (19) in order for there to be a damping ratio ξ minimal for ω =ωi it must fulfill the next equation: dξ dω = 1 2 (− α ωi + β). (20) after substitution ω =ωi into equation (20) we get α = ω21β (21) and then from equation (20) β = ξ1 ω1 . (22) in common practice, equation (17) is solved by direct integration using the newmark integration method. the discrete form of equation (17) already includes the influence of boundary conditions and the effect of loading in the form of a prescribed acceleration of displacements at the lower limit. 5. case exemplification the fem calculation results will be compared with the results of the analytical calculation published in [11]. the geometry of the model and the finite element mesh is shown in figure 6 and table 1 contains the parameters of the subsoil and the tunnel lining. the seismic loading has pga 2 ms−2 and the wave period is 0.01 s. all fem calculations were made by geo5 tunnel software [12]. for the sake of clarity the representative extreme values obtained for ϕ = φ4 have been set – see table 2. for illustration the resulting moment in the lining is also shown in figure 7. the two approaches in terms of the absolute percentage of the difference in the last column of table 2 can be mutually assessed. it is evident that in the case of a homogeneous mass and a circular excavation both methods give comparable results. the exception is the value of normal force in the case without a considering of the interaction between the lining and the soil (effect of the contact stress). if we take into account the value of contact stress in the equation (18), the resulting values ∆dff a ∆dlin of normal force for both methods are almost identical. type unit value unit loading γmean 0.0105 [-] soil gm = ghom 8.65 [mpa] νm 0.3 [pa] ρm 2000 [kg/m3] lining γmean 0.0105 [-] νl 0.3 [-] gl 12.5 [gpa] d 6 [m] h 0.3 [m] i 0.00225 [m4] table 1. input parameters of the numerical model figure 6. numerical model. 6. conclusions the paper is focused on the evaluation of the seismic effects on the tunnel lining by free-field analysis and pseudo static calculation using fem. generally, both calculation methods give comparable results in the case of a homogeneous soil environment and ignoring soil structure interaction. however, from a practical point of view, the fem combined with the 1d freefield analysis yields clear advantages: • it allows a possibility to derive the load for any time of the earthquake (measured or synthetically generated accelerograms). • it analyses any shape of excavation (lining). • it takes into account the potentially non-linear subsoil response. • it take into consideration the layered rock/soil mass depending on the "real" material parameters. another indisputable advantage is taking into account the soil structure interaction whose neglect in the case of an analytical method may lead to a significant overestimation of the normal force. such a non-realistic value of the normal force leads to a lower degree of reinforcement in the effects of the bending moment for the considered cross section. 41 jan pruška acta polytechnica ctu proceedings variable fem analytical difference [%] ∆ dff [mm] 31.3 31.5 0.6 ∆dlin [mm] 31.0 36.1 14.1 m [knm/m] –367.6 –422.8 13.1 v [kn/m] 244.5 281.0 13.0 n [kn/m] –380.8 –562.0 32.2 n∗ [kn/m] –380.8 –390.1∗ 2.4∗ * the influence of cont. nor. stresses was taken into the calc. table 2. output parameters. figure 7. moment in the lining. list of symbols a earthquake acceleration an ratio between the stiffness of lining and massive az,max pick acceleration at the depth of the tunnel d diameter of the excavation g gravitational acceleration h thickness of the lining k ratio of pgv to pga m1 weight of soil mass u actual displacement vector c ground motion at depth z c damping matrices cs apparent propagation velocity of s-wave d tunnel radius el young modulus of the lining f vector of nodal loading g shear modulus of soil/rock gm elasticity modulus of the massive i moment of inertia of the lining k stiffness matrices kv seismic coefficient in vertical direction kh seismic coefficient in horizontal direction m mass matrices q weight of structure pga peak ground acceleration r reduction factor of depth influence r radius of excavation rn lining – massive racking ratio under normal loading s coefficient based on the types of elastic response spectra suggested in eurocode 8 ss short period spectral acceleration vs peak ground velocity α parameter that determines the influence of mass in the damping of systems αm damping proportional to displacements β parameter that determines the influence of the stiffness in the damping of the system βk damping proportional strain rate γmean average shear strain of the model ηc integrate response coefficient related to the tunnel depth, soil property and project significance νm poisson number of the massive νl poisson number of the lining σv poisson number of the lining ω angular frequency ω1 angular frequency of the first natural mode ωi angular frequency of i-th mode acknowledgements the article was produced with financial support from the competence centres program of the technology agency of the czech republic (ta cr), project no. te01020168 centre for effective and sustainable transport infrastructure (cesti). references [1] j.-n. wang. seismic design of tunnel – a simple state of the art design approach. parsons, brinckerhoff, quade & douglas, inc. 1993. monograph 7. [2] m. s. power, d. rosidi, j. kaneshiro. strawman: screening, evaluation, and retrofit design of tunnels. national center for earthquake engineering research, buffalo, new york. report draft. vol. 3. [3] j. penzien, c. wu. stresses in linings of bored tunnels. international journal of earth-quake engineering and structural dynamics 3(27):283–300, 1998. [4] y. m. a. hashash, j. j. hook, b. schmidt, j. i.-c. yao. seismic design and analysis of under-ground structures. tunnelling and underground space technology (16):247–293, 2001. [5] e. bilotta, g. lanzano, g. russo, et al. pseudo-static and dynamic analyses of tunnels in transversal and longitudinal directions. in proceedings of 4th international conference on earthquake geotechnical engineering. icege, thessaloniki, 2007. 42 vol. 23/2019 underground structures subjected to seismic loads [6] m. corigliano, l. scandella, c. lai, r. paolucci. seismic analysis of deep tunnels in near fault conditions: a case study in southern italy. bulletin of earthquake engineering 9(4):975–995, 2011. [7] g. kouretzis, k. andrianopoulos, s. sloan, j. carter. analysis of circular tunnels due to seismic p-wave propagation, with emphasis on unreinforced concrete liners. computers and geotechnics 55:187–194, 2014. [8] aashto. technical manual for design and construction of road tunnels – civil elements. first edition. american association of state highway transportation officials, 2010. [9] c. hunk, j. monses, n. munfah, j. wisniewski. technical manual for design and construction of road tunnels – civil elements. first edition. federal highway administration, 2009. report fhwa-nhi-10-034. [10] g. lanzo, a. pagliaroli, b. d’elia. influenza della modellazione di rayleigh dello smorzamento viscoso nelle analisi di risposta sismica locale. in proceedings of 11th national conference l’ingegneria sismica in italia. 2004. [11] d. kučera. posouzení geotechnické konstrukce na seismické zatížení. diplomová práce. čvut v praze, fakulta stavební, 2017. in czech. [12] geo 5 fem. online, available from: https: //www.fine.cz/geotechnicky-software/mkp-tunel/. 43 https://www.fine.cz/geotechnicky-software/mkp-tunel/ https://www.fine.cz/geotechnicky-software/mkp-tunel/ acta polytechnica ctu proceedings 23:33–38, 2019 1 introduction 2 analytical method 2.1 seismic coefficient method 2.2 free – field deformation approach 2.3 soil-structure interaction approach 3 pseudo static fem 4 full dynamic calculations using fem 5 case exemplification 6 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0070 acta polytechnica ctu proceedings 11:70–73, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app the detailed priority of international freight expresses in the overlapping section of rfc 7 and rfc 9 kolín – choceň přemysl šrámek∗, tatiana molková university of pardubice, jan perner transport faculty, department of transport technology and control, studentská 95, 532 10 pardubice, czech republic ∗ corresponding author: premysl.sramek@student.upce.cz abstract. the main goal of this article is to research the optimal priority of international freight expresses, especially on european freight corridors. this article shows this problem in the context in the overlapping section of rfc 7 and rfc 9 kolín – choceň. there is solved changing priority of freight expresses in detail, whereas other train priorities are the same. keywords: capacity, freight expresses, freight corridors, timetable stability. 1. introduction one of the aims of european transport policy is to redirect capacity of road freight traffic to other modes of transport, where rail transport is an interesting and environmentally friendly alternative, especially in terms of speed, availability and amount of transported cargo. on the other hand, by customers is required delivery time is guaranteed by carriers, what is unfortunately not always possible, especially due to high utilization of european rail infrastructure. the solution is the construction of new lines, increasing the capacity of existing lines through construction and reconstruction measures or through operational measures. among the operational measures may be included alternative routing of trains as well as changing priorities of a particular type of train, e.g. train category fex (freight express) [1]. the aim of this article is to find the optimal priority of international freight expresses to satisfy the customers required delivery times due to optimization of train priority. 2. materials and methods there are mentioned some information about european freight corridors rfc and the simulation method with using of simulation tool simut (average delay increment (adi) calculation). 2.1. rail freight corridors (rfc) on the basis of the european parliament and the eu council no. 913/2010 for competitive freight and no. 1316/2013, which is created the connecting europe facility (cef), is the gradual establishment of rail freight corridors rfc (rail freight corridors). the main aim of the operation of these corridors is to strengthen the competitiveness of railways synergies between rail systems and harmonization of allocation interstate freight routes by national infrastructure managers [2]. czech republic, respectively czech railway infrastructure administration (cria), as the infrastructure manager and allocator of capacity, is a member of rfc 5 (baltic – adriatic), rfc 7 (orient / eastmediterranean), rfc 8 (north sea – baltic) and rfc 9 (czechoslovak corridor). each corridor is operated on the basis of the corridor information document (cid); each corridor has a single point of contact (c oss) and allows carriers to apply international prearranged train paths facilitated freight via c-oss. the request must be filed by information system rne pcs and must include the cross-border section. this paper deals with the model study of common section of rfc 7 and rfc 9 kolín choceň. 2.1.1. rfc 7 rfc 7 – the orient corridor runs from central europe to easternand southern-europe connecting 7 member states – czech republic, austria, slovak republic, hungary, romania, bulgaria and greece. the total length of main lines is approx. 3900 km and the length of alternative and connecting lines is almost 2500 km altogether. most limiting factors of rfc 7 are low capacity, speed limit, limited length of trains, limited axle load, not electrified sections and lack of adequate safety equipment (signalling track circuits with 25 hz frequency, etcs, gsm-r, etc.). the railway infrastructure managers and capacity allocation companies responsible for establishing and running rfc7 are committed to offer reliable, high-quality, competitive transport services in order to increase the market demand, to operate the infrastructure costeffectively on the long run through harmonization of technical and procedural conditions and to facilitate the environmentally sustainable development of the european economy and the achievement of a better quality of life for its people (3). in capacity analysis, published in implementation plan of rfc 7 [3], there are found lines with capacity utilization higher than 90 %. the longest sections with this high capacity 70 http://dx.doi.org/10.14311/app.2017.11.0070 http://ojs.cvut.cz/ojs/index.php/app vol. 11/2017 the detailed priority of international freight expresses utilization are situated in the czech republic, concretely sections poříčany – pardubice (65 km) and choceň – česká třebová (25 km). 2.1.2. frc 9 rfc 9 – the czech-slovak rail freight corridor (cs corridor) runs from prague to čierna nad tisou (slovak-ukrainian border) connecting 2 member states – czech republic and slovak republic. the total length of main lines is 972 km and the length of alternative and connecting lines is 276 km altogether. most limiting factors of rfc 9 are the same as in rfc 7. in capacity analysis, published in implementation plan of rfc9 [4], there are found lines with capacity utilization higher than 90 %. 2.2. the simulation method in this part is said something about the simulation method in context of the transit of international freight express trains (fex) in the overlapping section of rfc 7 and rfc 9 kolín – choceň. in this section there are a total amount of 10 railway stations, in which it is possible overtaking trains (double track line). interstation sections are divided by an automatic block into track sections; all crossing safety devices are equipped with gates. for simulation was used simulation tool simut, developed by pavel krýže, phd. from cria. the simulation tool simut is developed in visual basic, therefore it can run on every pc with microsoft office. it has to be input the option of a simulation, then railway stations and their shortcuts, station tracks for each railway station, line tracks, connection of station and line tracks, length of interstation departments and amount of line departments, type of train for priority, type of train for each number of a train and the path of each train. the simulation program simut can put new paths in a daily timetable with the solution for arisen path conflicts [1]. it was used daily timetable 2016, which included the amount of 384 trains. the similar topic was solved in (1), but not so exactly, in the evaluation there were used for fex only the importance values 0.5; 1.7; 1.9 and there was created periodic timetable. in the table 1 are displayed the importance values of trains, which remain the same, and the importance value of fex is increasing from 0.5 to 2.0 through step 0.1. therefore, it was solved 16 simulations, every simulation had 365 simulation runs (whole timetable period). about train types in table 1 – en means euronight train, sc is supercity train, ec – eurocity train, ic – intercity train, r is fast train, sp – speed-up passenger train, os – stop passenger train. the remaining train types are types of railway cargo transport – nex is here fex and its importance value is variable (the research object). pn is normal cargo train and mn means slow cargo trains servicing e. g. sidings. within the simulation program was established the average delay increment (adi). the average delay figure 1. adil in the relation with fex importance value. increment was calculated by dividing the difference between total output and total input delay and the total number of trains. this indicator was calculated as an ongoing basis for each simulation run, so the total for the entire graph (all simulation runs). the indicator was also calculated for different types of transport, i.e. for long-distance passenger transport (adil), regional passenger transport (adir) and freight transport (adif). as part of the simulation was set for all simulation runs random entry delay based on the exponential probability distribution. there were solved conflicts of station tracks, freight trains were allowed to ride before their schedule time (in the case of free capacity) [1]. 3. results and discussion in the table 2 there are displayed timetable stability indicators (adi, adil, adir, adif) in the relation with changing importance value of fex. all timetable stability indicators have unit min/train, it means every train of the category should get this delay in minutes. the timetable stability indicators are in the relation with fex importance value compared on the figure 1 (adil), figure 2 (adir), figure 3 (adif) and figure 4 (adi). adil is with increasing fex importance value just always increasing and it is logical – cargo transport has more and more preference. in the case, when fex importance value is the same as the ex importance value, it means 2.0, is adil strongly rising to 9.31 min/train. against fex important value 1.9 (4.39 min/train) is twofold value – the result is fex can’t have higher or the same priority as ex. on the figure 1 there is created the linking curve with the equation and reliability. for adil has the highest reliability exponential linking curve. on the figure 2 there is displayed adir – in this case is evaluation very difficult. the linking curve reliability is the best for polynomial linking curve. over 1.5 fex importance value is adir affirmative, the best compromise is for fex importance value 1.3 – adir is – 0.27 min/train. the result is fex should 71 přemysl šrámek, tatiana molková acta polytechnica ctu proceedings train type en sc ec ic ex r sp os sv nex pn mn lv importance value 2.5 1.8 3.0 2.8 2.0 1.8 1.5 1.0 0.3 var. 0.1 0.03 0.08 table 1. the train importance values. fex importance value adil adir adif adi 0.5 1.04 -0.54 1.19 0.84 0.6 1.19 -0.61 1.63 1.07 0.7 1.36 -0.37 1.21 1.02 0.8 1.42 -0.30 1.12 1.02 0.9 1.47 -0.51 1.31 1.09 1.0 1.88 -0.31 0.46 0.97 1.1 2.02 -0.08 0.31 1.01 1.2 2.22 -0.12 0.39 1.12 1.3 2.60 -0.27 -0.38 0.96 1.4 2.50 -0.06 0.79 1.42 1.5 3.05 0.05 0.25 1.47 1.6 3.08 0.01 0.57 1.60 1.7 3.00 0.06 -0.17 1.28 1.8 4.46 0.31 -0.36 1.89 1.9 4.39 0.01 -0.72 1.67 2.0 9.31 0.53 -1.44 3.66 table 2. timetable stability indicators in the relation of changing fex importance value. figure 2. adir in the relation with fex importance value. have higher priority than regional passenger trains, especially than stop passenger trains. the linking curve reliability for adif is the best for polynomial linking curve. from 1.7 fex importance value is adif always negative and decreasing, but negative is for 1.3 fex importance value, too. the result is here not so clear. on the figure 4 there is displayed average delay increment for whole timetable. in every case is the timetable unstable, adi is affirmative, the output delay is higher than the input delay. the linking curve has really high reliability for polynomial. for 2.0 fex importance value is adi getting really high due to increasing adil. less than 1 min/train is adi only in 3 cases – for fex importance values 0.5, 1.0 and 1.3. the best adi is for 0.5 (0.84 min/train), but figure 3. adif in the relation with fex importance value. figure 4. adi in the relation with fex importance value. it is in contrast with higher fex priority than regional passenger trains priority. the second best adi is for 1.3 (0.96 min/train), adil is 2.60 min/train, adir is – 0.27 min/train and adif is – 0.38 min/train. according to previous results the total result for best possible timetable stability is to set the fex importance value for chosen timetable to 1.3. and it corresponds the rules of prescription cria d1 [5], where the fex trains have importance between stop and speed-up passenger trains (1.0 – 1.5). 4. conclusions it was checked in detail the relation between the importance of fex trains and timetable stability. there were made 16 simulation variants in simulation program simut and then compared results. the research result is clear – for timetable 2016 in the overlapping section rfc 7 and rfc 9 kolín – choceň is necessary for best possible timetable stability to set the fex importance value (priority) to 1.3, it means between stop passenger trains priority and speed-up passenger trains priority. it corresponds in the czech republic 72 vol. 11/2017 the detailed priority of international freight expresses the contemporary prescription statement – it suffices only to keep it. acknowledgements this paper is supported by sgs. references [1] p.šrámek, t.molková. the priority of international freight expresses in the overlapping section of rfc 7 and rfc 9 kolín – česká třebová. studentská vědecká konference interoperabilita v železniční dopravě iricon 2016 – sborník, 2016. isbn 978-80-01-05939-5. [2] r.čech, m.šlachtová. železniční nákladní koridory rfc. železniční magazín czech raildays, m-presse plus, s. r. o., 2015. issn 1212-1851. [3] rfc 7 orient corridor, implementation plan of rail freight corridor 7 orient corridor [online], 2015. [cit. 2015-20-11]. [4] rfc 9 cs corridor, implementation plan of rail freight corridor 9 [online], 2013. [cit. 2015-23-11]. [5] sždc. d1 dopravní a návěstní předpis. sždc, s. o., 2013. 73 acta polytechnica ctu proceedings 11:70–73, 2017 1 introduction 2 materials and methods 2.1 rail freight corridors (rfc) 2.1.1 rfc 7 2.1.2 frc 9 2.2 the simulation method 3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.25.0025 acta polytechnica ctu proceedings 25:25–31, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app numerical modelling of wave shapes during shpb measurement radim dvořák∗, petr koudelka, tomáš fíla czech technical university in prague, faculty of transportation sciences, konviktská 20, 110 00 prague, czech republic ∗ corresponding author: dvorara9@fd.cvut.cz abstract. the paper aims at the numerical simulation of the wave propagation in compressive split hopkinson pressure bar (shpb) experiment. the paper deals with principles of shpb measurement, optimisation of a numerical model and techniques of pulse shaping. the parametric model of the typical shpb configuration developed for ls-dyna environment is introduced and optimised (in terms of element size and distribution) using the sensitivity study. then, a parametric analysis of a geometric properties of the pulse shaper is carried out to reveal their influence on a shape of the incident pulse. the analysis is algorithmized including the preand post-processing routines to enable automated processing of numerical results and comparison with the experimental data. results of the parametric analysis and the influence of geometric properties of the pulse shaper (diameter, length) on the incident wave are demonstrated. keywords: shpb, pulse shaping, finite element method, explicit dynamics. 1. introduction since material parameters are dependent also on strain rate, many experimental methodologies allowing to investigate material behavior during dynamic loading like fast impacts and explosions were developed. for example, yield strength of most of metals (e.g. copper) and alloys (including aluminium based alloys, steel, etc.) tends to increase with higher strain rates. the physics of this phenomena can be explained, for example, in a way that particles of material are affected by micro-inertia effects and do not have enough time to move and to cause yielding during high strain rates. however, quantification of this behavior is difficult. first experiments were performed by hopkinson [1], while an important improvement was proposed by kolsky [2] introducing the common split hopkinson pressure bar (shpb) apparatus [3]. shpb is an experimental device, which allows to measure the stress-strain diagram in dynamic compression, where the strain-rate and maximum compressive strain can be regulated within physical boundaries given by performance of the shpb apparatus and the investigated samples. in a typical configuration, the shpb consists of the impacting striker bar, the input incident bar, the output transmission bar, and the specimen placed between the incident and transmission bars (see figure 1). all the bars are manufactured from the same material and have the same diameter, while the incident and transmission bar also have the same length selected according to parameters of the experiment and properties of the investigated specimen. typical instrumentation of the experiment consists of strain-gauges mounted on the incident and transmission bars, through-beam sensors for measurement of the striker impact velocity, and a high-speed camera for optical inspection of the experiment. the accelerated striker impacts the incident bar forming a stress-wave propagating through the incident bar until the bar-specimen interface is reached. on the specimen boundary, the part of the incident wave is reflected and part is transmitted through the specimen causing its deformation. the transmitted part of the initial pulse is recorded by the transmission bar strain-gauge. during the experiment, incident, reflected and transmitted pulses are recorded with strain-gauges (see figure 2). the recorded strain-gauge signals and known properties of the experimental setup (e.g. bar dimensions, young’s modulus, density, wave propagation velocity etc.) are used for the evaluation of the stress-strain diagram of the specimen for the given strain-rate. for valid experiment, two crucial requirements have to be fulfilled: i. satisfactory dynamic stress equilibrium and ii. approximately constant strain-rate. in the specimen, the wave is also reflected on the interface of the specimen with the transmission bar. therefore, a defined number of reflections has to pass through the specimen before the dynamic stress equilibrium is reached (see figure 3). the dynamic equilibrium condition in the experiment is crucial as the standard mathematical methods for the evaluation of the specimen response are valid only in the equilibrium state. ramp-in effects (also known as ringing effects a transient state caused by sudden incidence of a sharp pulse) and spurious oscillations of the specimen, negatively affecting dynamic equilibrium, can be significantly reduced by a pulse-shaping technique. a pulse shaper is usually a small cylinder of a soft material 25 http://dx.doi.org/10.14311/app.2019.25.0025 http://ojs.cvut.cz/ojs/index.php/app r. dvořák, p. koudelka, t. fíla acta polytechnica ctu proceedings figure 1. shpb apparatus figure 2. three measured signals figure 3. space-time diagram of the wave (soft copper, soft aluminium, paper etc.) with length in order of tenths or units of mm placed between the striker and the input bar. the pulse shaper causes slower raising of the incident wave and also attenuation of high frequency oscillations (see figure 4). as a result, the pulse shaper improves conditions of the experiment leading to good quality dynamic equilibrium and constant strain-rate during loading of the specimen [4–6]. in this paper, we concentrate on numerical aspects of simulations of the shpb experiments. finite element (fe) model of shpb was developed in ansys ls-dyna. a sensitivity analysis of the fe model of the bars was performed as an important part of this study to assess the influence of the selected meshing algorithm and element size on numerical results. then, the numerical model was calibrated according to experimental data. the second part of this contribution summarizes a parametric study of a pulse-shaper. length, diameter and impacting velocity were varied parameters and evaluation of their influence on shape of stress pulse was analyzed. figure 4. original and shaped wave 2. models and methods a numerical model of shpb apparatus consisting of striker bar, incident bar and the pulse shaper was created in ansys. the bars had the same diameter of 20 mm. no global constraints were prescribed to the numerical model. two-way surface to surface algorithm was used for contact boundary conditions between the individual elements of virtual shpb apparatus. two algorithms of meshing, implemented in ansys toolkit were compared (see figure 5). the direct volume meshing algorithm creates elements from within of the bars and causes generating very small elements on the edges of the bar. these small elements have negative influence on the computing time, because the time step size is calculated (and is proportional to) from the smallest element size. the direct meshing algorithm was compared with the sweeping algorithm for mesh generation, which chooses two faces that are topologically on the opposite sides of the body. these faces are called the source and target faces. algorithm meshes the source face with quadrilateral and triangular elements and then copies that mesh onto the target face. then it generates either hexahedral or wedge elements connecting the two faces and following the exterior topology of the body [7]. using this algorithm, larger elements were produced in the middle of the bar (see figure 5). the variation of element sizes and internal angles is apparently higher, when compared to directly generated mesh, but provides control over the achievable time-step via setting the mesh size on the edge of the bar. of course it is possible to create a regular mesh using meshing software specialized for geometries such as cylinders (ls-dyna/lspp), however, ansys was used as a pre-processor in this case for the ease of 26 vol. 25/2019 modelling of wave shapes during shpb measurement figure 5. outcome of the direct (top) and the sweeping method (bottom). automation, hence the mesh is a result of its meshing algorithms with only limited controlability. however, in our experience with simulating the shpb problems, the irregularity of the sweeped mesh produces negligible differences to ls-dyna/lspp cylinder meshes for most cases. after the sensitivity studies are concluded and critical factors identified, we use best quality approach for mesh generation. however, it was necessary to compare both mesh generation algorithms in terms of influence on the virtual stress-wave propagation. for this reason, stress wave was recorded on the impacted front of the bar and in the distance of two and four multiples of bar radii measured from the front of the bar (see figure 6). the stress-wave is almost identical for both models as can be seen in figure 6, in the distance equal to 2 diameters from the impact face. it is evident that from the distance of 2 diameters from the impacted face of the bar both models have the same stress wave evolution. this results proves the sweeping algorithm as a more suitable method for our task due to ability of indirectly control the time step using the size of the element. in this case, the direct meshing algorithm produced elements with five times shorter time step than sweeping algorithm. mentionable is the difference in the evolution of the stress at the striker-bar interface (0 r). for both cases automatic surface to surface contact algorithm figure 6. an evolution of the stress wave measured at the front of the bar (top), one diameter distance (middle) and two multiples of the diameter distance (bottom) was used. boundary conditions are identically the same for both models. this anomaly is probably the consequence of the contact algorithm failure at the edge of the cylinder, which is supposed to happen very occasionally. the algorithm probably could not recognize if the bars are in or out of the contact and was alternately connecting and disconnecting the coupling between interfaces. this phenomena should be investigated in further studies, however now is not essential for our purpose, that is to analyze pulse shaper influence on the stress wave. although both model results are from determined border of 4 radii comparable with theoretical assump27 r. dvořák, p. koudelka, t. fíla acta polytechnica ctu proceedings figure 7. comparison of the finest, the middle and the coarse mesh tions of a wave propagating phenomena, it is well known that fem can produce very unreliable results particularly in the analysis of stress waves propagation using explicit numerical methods. basically, smaller elements cause that higher frequencies pass through the material in the model. the chosen sweeping algorithm model had to be calibrated according to the experimental data. a series of void tests (impacts of the striker bar on the incident bar without a pulse shaper) using shpb was performed and the experimentally measured signals were compared with the signals from the fe model. unphysical behavior of the stress signal in the model with the rough mesh (40 elements along the circumference of the bar) can be seen in figure 7. small sized elements in the model allowed for propagation of high-frequencies that were not observed during the experiments. however, choosing between the middle sized (160 elements along the circumference) element model and the finest one (280 elements along the circumference) is not possible without comparing with experimental data. results of the models with the middle-sized and coarse mesh are shown in figure 8. it can be seen that results of the coarse mesh model showed the best agreement with the experimental data. on the basis of the results of the sensitivity study, a set of automated parametric simulations comprising ansys environment, ls prepost tool, ls-dyna solver, and matlab was peformed to assess influence of pulse-shaping on stress-waves. for the parametric study, the thickness and diameter of a cylindrical pulse-shaper were selected as the control variables. in the simulations, the diameter of pulse-shaper in the range of 5 − 8 mm and its thickness in the range of 0.5 − 1.5 mm were considered and the simulations were performed for the striker impact velocity in the range of 5 − 50 mm · s−1. 3. results in total, 130 simulations were peformed and the results were derived from the stress-wave shapes captured at the distance of four radii of the bars from its impacted figure 8. comparison of experiment data with the middle (top) and the rough (bottom) meshed model face. changes of the wave shape caused by different pulse-shaper geometries during the impact with velocity of 25 mm · s−1 can be seen in figure 9. the fe modelling was identified as a very efficient tool to investigate the phenomena. observed attributes of the wave were: (1.) the wave-length of the pulse, defined as the time interval between the first nonzero value of stress and the first decrease of stress bellow the value of 5 % of the maximum value achieved during measuring, (2.) the rise time of the wave, defined as the time interval between the first nonzero value of stress and achieving the first peak, (3.) the rate of oscillations in the first half of the pulse, defined as the sum spread from smoothed data in the time interval between the first nonzero stress value and half of the above defined pulse length. the length of the pulse slightly incrreases with longer pulse shaper (see figure 10). the fluctuations are smoothed. the time rise also slightly rises, however the effect of changing length is not as remarkable as using pulse shaper by itself. changing the diameter has an apparent influence on the attributes of the wave (see figure 11). with smaller diameters the wave is longer (1.2 ∼ 1.4 times), the time rise is also longer and oscillations get smoothed. these attributes of the pulse are ap28 vol. 25/2019 modelling of wave shapes during shpb measurement figure 9. influence of the pulse-shaper geometry figure 10. influence of the length of the ps propriate for achieving the stress equilibrium in the specimen. 3.1. the influence of the strain rate on the time rise for different geometries of the pulse shaper the results have a comparable trend except two cases (see figure 12) that can be identified as numerical error (probably during automatic postprocessing, another possibility is an imperfect processing of these two simulations e.g., due to not defined friction at the contact faces resulting in slipping of the pulse shaper and that negatively influenced the data). for diameters of 5 and 6 mm no distortion of results occurred, which means this mesh is capable to describe high frequencies generated by high-velocity impact (40 − 50 mm · s−1). the other curves have a comparable manner and two mentioned cases should be an error. it is evident that decreasing the diameter causes remarkable prolonging (from 2 to 4.5 times) of the rise time and increasing the length causes insignificant prolonging of the rise time. both effects are stronger during lower strain rates. figure 11. influence of the diameter of the ps 3.2. the influence of the strain rate on the length of the pulse for different geometries of the pulse shaper data for 7 and 8 mm pulse shaper are again nonnegligibly distorted and it is caused by the identical numerical error mentioned in the section above. moderate influence of the strain-rate is evident (see figure 13). lower strain rates cause prolonging (from 1.06 to 1.22 times) of the wave and this effect is more remarkable with decreasing diameter. length of the pulse shaper has a minimal influence, the remarkable cases are is 5 and 6 mm. 3.3. the influence of the strain rate on the oscillations of the pulse for different geometries of the pulse shaper it is evident, that pulse shapers with smaller diameters behave as a better filter (see figure 14). also the filtering of high frequencies is more effective during lower strain rates. the length of the pulse shaper has a negligible influence during slow impacts (5 − 30 mm · s−1). during high velocity impacts (40 − 50 mm · s−1) also the length has a remarkable influence. 4. conclusions the numerical model of the shpb apparatus was calibrated including the mesh sensitivity study. two types of mesh generation algorithms were analyzed and boundary where the stress wave has the similar shape for both models was identified. the border where the stress curve has similar shape for both models was established, size of elements was set up according to experimental data. a phenomenon that the mesh actually behaves as a filter of high frequencies was observed. following conclusions can be drawn based on the main parametric study: (1.) pulse-shapers with smaller diameter are more effective as filters of oscillations, well prolong the 29 r. dvořák, p. koudelka, t. fíla acta polytechnica ctu proceedings figure 12. influence of impacting velocity on a time rise (l length of the pulse shaper, d diameter of the pulse shaper figure 13. influence of impacting velocity on a pulse length (l length of the pulse shaper, d diameter of the pulse shaper figure 14. influence of impacting velocity on oscillations of the pulse (l length of the pulse shaper, d diameter of the pulse shaper 30 vol. 25/2019 modelling of wave shapes during shpb measurement time rise and the pulse length which is crucial for achieving stress equilibrium. (2.) the length of the pulse shaper has an insignificant influence. each mentioned phenomena is more remarkable during low strain rates. these conclusions are fully compatible with other studies [4–6] and with the experiments. list of symbols σ stress [mpa] εi incident pulse [–] εt transmitted pulse [–] εr reflected pulse [–] t time [ms] acknowledgements the research was supported by the czech science foundation (project no. 19-23675s). all the financial support is gratefully acknowledged. references [1] b. hopkinson. the scientific papers of bertram hopkinson. cambridge university press, 1921. [2] h. kolsky. an investigation of the mechanical properties of materials at very high rates of loading. proceedings of the physical society section b 62(11):676– 700, 1949. doi:10.1088/0370-1301/62/11/302. [3] w. w. chen, b. song. split hopkinson (kolsky) bar: design, testing and applications (mechanical engineering series). springer, 2010. [4] r. naghdabadi, m. ashrafi, j. arghavani. experimental and numerical investigation of pulse-shaped split hopkinson pressure bar test. materials science and engineering: a 539:285–293, 2012. doi:10.1016/j.msea.2012.01.095. [5] r. gerlach, s. k. sathianathan, c. siviour, n. petrinic. a novel method for pulse shaping of split hopkinson tensile bar signals. international journal of impact engineering 38(12):976–980, 2011. doi:10.1016/j.ijimpeng.2011.08.007. [6] a. b. shemirani, r. naghdabadi, m. ashrafi. experimental and numerical study on choosing proper pulse shapers for testing concrete specimens by split hopkinson pressure bar apparatus. construction and building materials 125:326–336, 2016. doi:10.1016/j.conbuildmat.2016.08.045. [7] j. hallquist. ls-dyna keyword users manual. livermore software technology corp, 2006. 31 https://doi.org/10.1088/0370-1301/62/11/302 https://doi.org/10.1016/j.msea.2012.01.095 https://doi.org/10.1016/j.ijimpeng.2011.08.007 https://doi.org/10.1016/j.conbuildmat.2016.08.045 acta polytechnica ctu proceedings 25:25–31, 2019 1 introduction 2 models and methods 3 results 3.1 the influence of the strain rate on the time rise for different geometries of the pulse shaper 3.2 the influence of the strain rate on the length of the pulse for different geometries of the pulse shaper 3.3 the influence of the strain rate on the oscillations of the pulse for different geometries of the pulse shaper 4 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.18.0020 acta polytechnica ctu proceedings 18:20–23, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app instrumentation of four-point bending test during 4d computed tomography daniel kytýřa, ∗, tomáš fílaa, petr koudelkaa, ivana kumpováa, michal vopálenskýa, leona vavrob, martin vavrob a czech academy of sciences, institute of theoretical and applied mechanics prosecká 809/76, 190 00 prague 9, czech republic b czech academy of sciences, institute of geonics studentská 1768, 708 00 ostrava-poruba, czech republic ∗ corresponding author: kytyr@itam.cas.cz abstract. high-resolution time-lapse micro-focus x-ray computed tomography is an effective method for investigation of deformation processes on volumetric basis including fracture propagation characteristics of non-homogeneous materials subjected to mechanical loading. this experimental method requires implementation of specifically designed loading devices to x-ray imaging setups. in case of bending tests, our background research showed that no commercial solution allowing for reliable investigation of so called fracture process zone in quasi-brittle materials is currently available. thus, this paper is focused on description of recently developed in-situ four-point bending loading device and its instrumentation for testing of quasi-brittle materials. proof of concept together with the pilot experiments were successfully performed in a ct scanner toratom. based on results of the pilot experiments, we demonstrate that crack development and propagation in a quasi-brittle material can be successfully observed in 3d using high resolution 4d micro-ct under loading. keywords: in-situ loading device, four-point bending, 4d micro-ct, crack propagation. 1. introduction proper characterization of damage processes in nonhomogeneous materials requires development of advanced experimental procedures and instrumentation of testing devices. one of currently investigated problems is evaluation of the so called fracture process zone (fpz) [1] together with crack propagation in quasi-brittle building materials [2], e.g. sandstones or man-made cementitious composites. for this purpose, notched samples are typically submitted to flexural loading in standard loading frames. in case of optical or thermographical measurement, basic information about crack propagation can be obtained, but information about deformation processes and fpz evolving within volume of the tested sample stays uncovered. for that reason, radiographical observation of the loading procedure capable of revealing microstructural variations in 3d is one of the most efficient methods [3]. to obtain relevant information about deformation and fracture behaviour of investigated quasi-brittle material, development of complex experimental procedure including design of a specialised loading device was necessary. we proposed a novel approach to four-point bending experiments during high-resolution micro-ct scans, which is based on vertical orientation of the investigated specimen, whose axis of rotations is identical with rotational axis of the ct scanner. this arrangement significantly reduces fluctuation of x-ray attenuation during the imaging and thanks to smaller source-object distance higher resolution of the reconstructed 3d images can be achieved during the experimental procedure. aside from technical details of the experimental device, proof of concept together with the preliminary results are presented in this paper. 2. fracture of quasi-brittle material fracture of quasi-brittle materials is accompanied by formation and evolution of an inelastic zone around the propagating crack tip. this area referred to as the fpz, which starts to appear at ≈ 75 % of ultimate stress is responsible for the non-linear propagation of the crack in quasi-brittle materials [4]. typical deformation response to the loading of the quasi-brittle material together with scheme of fpz is depicted in fig. 1. in this zone, the damage on several scale levels of material microstructure can be detected. since the fpz is a non-linear zone of damage in front of the macroscopic crack propagating through the material, its influence on overall damage resistance of quasi-brittle structural elements has to be treated with respect to ratio of fpz size to size of the assessed structural element. it has already been shown that size of the fpz may be comparable to the material’s basic constituents (e.g. the aggregates and voids) or it can be greater in extent by up to two orders of magnitude [5]. here, x-ray micro-ct measurement plays crucial role in detection of the crack tip and fpz propagation in the materials’ microstructure. 20 http://dx.doi.org/10.14311/app.2018.18.0020 http://ojs.cvut.cz/ojs/index.php/app vol. 18/2018 instrumentation of 4pb test during 4d ct crack tip fpz f o rc e displacement figure 1. typical f–d curve of a quasi-brittle material subjected to flexural loading (left), schematic propagation of fpz (right) 3. instrumentation since 2013, the modular system for x-ray based observation of time-dependent processes in wide range of structural and functional materials has been continually developed to fulfill recent challenges of experimental mechanics and to keep all technologies of the system up-to-date. the system consists of the patented toratom imaging device (european patent no. ep2835631) and various loading device for specific experimental purposes. 3.1. x-ray imaging device in the current configuration, the device consists of two modular perpendicular imaging chains, which enables dual-source/dual-energy radiographical and tomographical experiments. wide range of its applications is given by various x-ray sources and detectors currently available. the first xwt 160 tchr (x-ray worx, germany) 160 kev microfocus transmission tube with 1 µm focal spot at target power of 3 w can be used for detailed analysis of materials with complex microstructure. the second high power xwt 240 se (x-ray worx, germany) 280 w reflection tube allows to analyse physically large samples or samples of materials with high attenuation of x-rays. furthermore, it is possible to use different types of radiation imaging detectors scintillating large-area flat panel detector, or various semiconductor pixelated detectors of the medipix family [6, 7] with theoretically unlimited dynamic range, including the large area detector composed of 10 × 10 timepix 2 assembly [8], fast devices with maximum possible frame rate of 200 fps without compression in two-row/twocolumn configuration of timepix chips [9], and also spectroscopic detectors. for the analysis presented in this paper, a combination of the xwt 160 source and xrd1622 (perkinelmer, usa) scintillating large-area flat panel detector was used. 3.2. image correction procedure it is a well known fact that the raw radiograms cannot be usually used for a high-resolution ct reconstruction procedure with sufficient quality of the reconstructed 3d image as they may suffer from stationary noise caused by the inhomogeneous characteristics of individual pixels, inhomogeneity of the background brightness caused by the intensity profile of the x-ray beam, and abnormal response of individual pixels (so called bad pixels). these effects then form different types of artifacts in the 3d reconstruction. for their elimination, the flat field correction [10] with equalization of the images and bad pixel correction were applied. these corrections were performed using the torast in-house developed python software (for a usage example see [11]). all the functions are available in a batch mode, allowing a comfortable processing of whole sets of projection images. new features are currently being integrated, such as the compensation of focal spot movement, elimination of residual gradients in the flat-field corrected images). in the future, the software will be integrated with the control software of the toratom device to enable advanced acquisition methods. 3.3. loading device the concept for 4d micro-ct of the bending experiments was developed based on experience from measurements presented in [12, 13], where conventional three-point bending tests (see fig. 2 left) were employed. although it was possible to obtain quantification of the crack-tip propagation processes, the horizontal orientation of the sample perpendicular to the rotational axis of the ct scanner together with the loading supports covering parts of the radiograms caused limitations in terms of contrast and resolutions. loading rotation axis loading loading rotation axis four−point bending test three−point bending test analysis region analysis region figure 2. basic principles of the loading scenarios these disadvantages were overcome by employment of four-point bending setup, where the sample is oriented vertically with its longitudinal axis identical with the rotational axis of the ct scanner (see fig. 2 right). the newly developed device (czech national patent no. pv2018-28 pending) is composed of three main 21 d. kytýř, t. fíla, p. koudelka et al. acta polytechnica ctu proceedings figure 3. components of the 4pb loading device components: a pair of motorized loading units with integrated movable outer supports, a pair of a stationary inner supports, and the cylindrical load bearing frame housing the loaded specimen together with the components of the loading device. the design of the device in presented in fig. 3. the loading device is during the radiographically observed loading sequence oriented vertically and fixed on rotary stage of the ct scanner. to ensure high contrast in the acquired radiograms, the part of the loading frame located inbetween the inner supports was manufactured from carbon fiber composite (t700s carbon fibers, mtm57 series epoxy resin) with low attenuation of x-rays with the nominal thickness 1.95 mm. the outer movable supports are integrated with driving units with precision captive stepper linear actuators (23-2210, koco motion dings, usa) and precision linear guideways (mgw12, hiwin, japan). both driving units are independent and equipped with linear encoders with resolution 1 µm (lm10, renishaw, uk) for highly precise positioning. each driving unit is also equipped with individual load cell (lcm300, futek, usa) for force measurements on the outer supports. the maximum load capacity of the device is 1500 n per single support. position accuracy and repeatability in single-direction loading is better than 10 µm with sensitivity better than 1 µm. 3.4. experimental procedure the batches of specimens in shape of a cylinder with diameter d = 36 mm and length l = 195 mm were drilled out from the godula sandstone, kocbeře sandstone, and mšené sandstone that all represent traditional building and decorative natural stones widely used in masonry for about a thousand years [14]. the core drilling was carried out parallel to the sandstone bedding planes. in central part of the sample, a chevron notch was prepared using a circular diamond blade. the specimen was subjected to four-point bending loading sequence performed as a force-driven experiment to ensure uniform load distribution on the supports (outer supports span lout = 180 mm, inner supports span lin = 150 mm). experiments were controlled by the modified linuxcnc software running on a real-time kernel [15]. during the experiment, displacement was interrupted at six discrete loading steps, where radiographical imaging was performed. in the loading sequence, one load-step was performed during the material’s hardening phase, one load-step at the ultimate-stress point (peak force reached was approx. 325 n), and four load-steps were performed during the post-peak softening phase. the reconstructed image data from the notch region with resolution better than 20 µm showing evolution of the crack and the related changes in microstructure are depicted in the fig. 4. figure 4. visualization of damaged sample in three perpendicular planes based on the reconstructed microct data corresponding to displacement 320 µm 4. conclusions proof of concept based on the pilot experiments was successfully performed in a micro-ct scanner. based on these experiments, it was shown that crack development and propagation in a quasi-brittle material can be observed in 3d using in-situ loading procedure and high resolution 4d micro-ct imaging. arrangement of the new experimental setup together with higher stiffness of the loading device allows to interrupt the loading procedure at any point of the material’s f–d curve without the risk of sudden collapse of the sample. concept of vertically oriented sample also allows to increase signal to noise ratio in the acquired projections to achieve higher contrast in the reconstructed 3d images. 22 vol. 18/2018 instrumentation of 4pb test during 4d ct acknowledgements the research has been supported by the european regional development fund in frame of the project com3dxct (atcz38) in the interreg v-a austria czech republic programme and by operational programme research, development and education in project inafym engineering applications of microworld physics (cz.02.1.01/0.0/0.0/16_019/0000766). references [1] j. klon, v. vesely. modelling of size and shape of damage zone in quasi-brittle notched specimens – analytical approach based on fracture-mechanical evaluation of loading curves. frattura ed integrita strutturale 11(39):17–28, 2017. doi:10.3221/igf-esis.39.03. [2] v. vesely, z. kersner, i. merta. quasi-brittle behaviour of composites as a key to generalized understanding of material structure. procedia engineering 190:126–133, 2017. doi:10.1016/j.proeng.2017.05.317. [3] l. skarzynski, j. suchorzewski. mechanical and fracture properties of concrete reinforced with recycled and industrial steel fibers using digital image correlation technique and x-ray micro computed tomography. construction and building materials 183:283–299, 2018. doi:10.1016/j.conbuildmat.2018.06.182. [4] t. anderson. fracture mechanics: fundamentals and applications. third edition. crr press, 2004. [5] v. vesely, z. kersner, j. nemecek, et al. estimation of fracture process zone extent in cementitious composites. chemicke listy 104(15 spec):s382–s385, 2010. [6] j. jakubek. data processing and image reconstruction methods for pixel detectors. nuclear instruments and methods in physics research section a: accelerators, spectrometers, detectors and associated equipment 576(1):223–234, 2007. doi:http://dx.doi.org/10.1016/j.nima.2007.01.157. [7] e. gimenez, v. astromskas, i. horswell, et al. development of a schottky cdte medipix3rx hybrid photon counting detector with spatial and energy resolving capabilities. nuclear instruments and methods in physics research, section a: accelerators, spectrometers, detectors and associated equipment 824:101–103, 2016. doi:10.1016/j.nima.2015.10.092. [8] i. kumpova, d. vavrik, t. fila, et al. high resolution micro-ct of low attenuating organic materials using large area photon-counting detector. journal of instrumentation 11(2), 2016. doi:10.1088/1748-0221/11/02/c02003. [9] d. kytyr, p. zlamal, p. koudelka, et al. deformation analysis of gellan-gum based bone scaffold using on-the-fly tomography. materials and design 134:400–417, 2017. doi:10.1016/j.matdes.2017.08.036. [10] j. rowlands, y. j. flat panel detectors for digital radiography. in handbook of medical imaging, vol. 1: physics and psychophysics. spie press, 2000. [11] m. vopalensky, d. vavrik, i. kumpova. optimization of the x-ray tube voltage with respect to the dynamical resolution in radiography and tomography. ndtnet journal 2018-02, 2018. [12] l. vavro, k. soucek, d. kytyr, et al. visualization of the evolution of the fracture process zone in sandstone by transmission computed radiography. procedia engineering 191:689–696, 2017. doi:10.1016/j.proeng.2017.05.233. [13] i. kumpova, m. vopalensky, t. fila, et al. on-the-fly fast x-ray tomography inspection of the quasi-brittle material three point bending test. 2016 ieee nuclear science symposium, medical imaging conference and room-temperature semiconductor detector workshop 2017-january, 2017. doi:10.1109/nssmic.2016.8069950. [14] m. vavro, l. vavro, p. martinec, k. soucek. properties, durability and use of glauconitic godula sandstones: a relatively less known building stone of the czech republic and poland. environmental earth sciences 75(22), 2016. doi:10.1007/s12665-016-6248-3. [15] v. rada, t. fila, p. zlamal, et al. multi-channel control system for in-situ laboratory loading devices. acta polytechnica proceedings ctu proceedings 18:15–19 2018. doi:10.14311/app.2018.18.0015. 23 http://dx.doi.org/10.3221/igf-esis.39.03 http://dx.doi.org/10.1016/j.proeng.2017.05.317 http://dx.doi.org/10.1016/j.conbuildmat.2018.06.182 http://dx.doi.org/http://dx.doi.org/10.1016/j.nima.2007.01.157 http://dx.doi.org/10.1016/j.nima.2015.10.092 http://dx.doi.org/10.1088/1748-0221/11/02/c02003 http://dx.doi.org/10.1016/j.matdes.2017.08.036 http://dx.doi.org/10.1016/j.proeng.2017.05.233 http://dx.doi.org/10.1109/nssmic.2016.8069950 http://dx.doi.org/10.1007/s12665-016-6248-3 http://dx.doi.org/10.14311/app.2018.18.0015 acta polytechnica ctu proceedings 18:20–23, 2018 1 introduction 2 fracture of quasi-brittle material 3 instrumentation 3.1 x-ray imaging device 3.2 image correction procedure 3.3 loading device 3.4 experimental procedure 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0020 acta polytechnica ctu proceedings 13:20–28, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of methods used for seismic analysis of structures petr čada∗, jiří máca czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: petr.cada@fsv.cvut.cz abstract. this paper investigates effects of the seismic load to a structure. the article describes main methods of the definition and practical application of the seismic load based on the standard eurocode 8. there was made a comparison of all methods using the same structure. a simple two-storeyed concrete 2d-frame with fixed joints was chosen. a one another model with rigid beams for some calculations was defined. the second model can be used for hand-calculations as a cantilever with two masses. the paper describes main dynamic properties of the chosen structure. seismic load was defined by lateral force method, modal response spectrum, non-linear time-history analysis and pushover analysis. the time-history analysis is represented by accelerograms. there were made linear and non-linear calculations. keywords: seismic response, lateral force method, response spectrum, time-history analysis, pushover analysis. 1. introduction for a calculation of the seismic response, the linear state of stress is commonly used. by complicated cases or by higher importance of the structure is recommended to use some non-linear method. in eurocode 8 (ec8) [1] these four methods of analysis are defined: • lateral force method, • modal response spectrum analysis, • non-linear time-history (dynamic) analysis, • non-linear static (pushover) analysis. for the better results of the behaviour of the structure, there is necessary to use one of the non-linear methods. the ductility is the most important property for a non-linear calculation. the ductility describes the behaviour of the plastic deformations, which cause irreversible deformations and this fact can be used for the reduction of the seismic load. calculations in this article describe the possibilities of the ec8 [1]. for this purpose, a two-storeyed concrete frame with a uniformly distributed load was used. this simple structure was chosen because of the possibility to applicate all methods. only a simple model can fulfil all conditions of all methods. lateral force method is a static and linear (linear behaviour of the material) calculation using horizontal forces as a seismic load. by this method a high frequency cannot have the decisive influence. the base shear force can be defined for each direction. distribution of horizontal forces is linear increasing or responds to the eigenvalues. the each storey must be rigid in its plane [1]. modal response spectrum analysis is a dynamic method with the same geometrical rules as by the lateral force method. for the horizontal displacement, all decisive eigenmodes should be used, but at least 90 % of the total mass. responses from different directions can be combined by the srss (square root of the sum of the squares) or cqc (complete quadric combination) method. this is possible only when eigenmodes are mutually independent [1]. worked examples for lateral force method and modal response spectrum are shown in [2]. non-linear time-history (dynamic) analysis is based on direct numerical integration of the differential equation of motion. motion of the base is represented by the accelerogram [1]. this method is described in [3, 4] as well. non-linear static (pushover) analysis is the last type of the seismic calculations from the ec8 [1]. in this standard, the most used version – n2 method – is used. the name was derived from “nonlinear” for “n” and “2” for two mathematical models. method was developed by prof. peter fajfar. three steps of analysis are defined. the first step is used for determination of the stiffness, strength and ductility of the structure. the first mathematical model – mdof (multiple degree of freedom) system – with linear increasing horizontal load is used. for the structure, an equivalent sdof (single degree of freedom) system in the second step must be defined. the behaviour must be same as by the mdof system. the ratio between base shear force and the top-displacement is the base for the non-linear characteristics. the maximal displacement is calculated in the third step by the sdof system. the pushover analysis is regularly used for the first eigenmode (the first eigenmode must 20 http://dx.doi.org/10.14311/app.2017.13.0020 http://ojs.cvut.cz/ojs/index.php/app vol. 13/2017 comparison of methods used for seismic analysis of structures be decisive), but theoretically is it possible to use this method for any eigenmod [5]. for example [6–9] are dealing with this issue as well. 2. structure and load 2.1. model of the structure and material parameters for sample-calculations of the response to the seismic load, a simple two-storeyed concrete frame with fixed joints was chosen. the width is 6 m and the height is 4 m for one storey. the total height is 8 m. the frame consists of two cross-sections. columns are square cross-section of the dimension 40×40 cm. the reinforcement is distributed symmetrical to all sides and the bar diameter is ø20. cross-beams are rectangular cross-section of the dimension 40×60 cm and the reinforcement is placed only at the bottom and on the top of cross-section. it is used 3ø20 for one side. the shear-reinforcement is not considered. figure 1. 2d-model of the construction. figure 2. cross-sections. concrete strength class is c30/37 and the reinforcement strength class is b500b. main material parameters for linear and non-linear behaviour were defined according to [10], chapter 3.1 and 3.2. the most important parameters for the non-linear analysis of the concrete are compressive strength fck = 30 mpa, tension strength (5% fractile) fctk0,05 = 2.0 mpa and design strain εcd = 0.077 ‰. for the reinforcement are the most important yield strength fyk =ftk = 500 mpa and strain εyk = 2.5 ‰. 2.2. load 2.2.1. vertical load the frame is loaded by the self-weight (6 kn/m for the cross-beam and 4 kn/m for the column) and a uniformly distributed load 10 kn/m, which represents fittings and imposed loads together. for the simplification, only one load for one combination was defined. safety factors are not considered as well. figure 3. vertical load. 2.2.2. seismic load seismic load was represented by horizontal elastic response spectrum, which is defined in [1], chapter 3.2.2.2. initial conditions were chosen randomly, because the structure is fictional. the design peak ground acceleration (pga) is 0.8 m/s, that is 0.0815g. the response spectrum is type 1, which corresponds in the czech republic to the moravia and silesia. the frame is based in the ground type b, which corresponds to the very dense sand, gravel, or very stiff clay, with the average shear wave velocity vs,30 = 360 m/s. the importance class of the structure is class ii, which describes ordinary buildings with the importance factor γ,i = 1.0. because the comparison has to follow the original response of the structure, it was necessary to use the 21 petr čada, jiří máca acta polytechnica ctu proceedings figure 4. horizontal elastic response spectrum. behaviour factor q = 1. with the behaviour factor q = 1, it cannot to occur to an undesirable dissipation of the energy and to the change of the result. the effective damping (ξ) follows the recommendation of the standard [1] and it is ξ = 5 %, which implies that the damping correction factor is equal to 1 (η = 1). 2.3. software for following calculations was necessary to use some software: • for the definition of the accelerogram, a programm seismoartif [11] was used. seismoartif [11] is made by the company seismosoft (pavia, italy). this programm allows a conversion of the response spectrum to the one or more accelerograms. accelerograms are used for the time-history analysis. • for calculations with accelerograms, a programm seismostruct [12] was used. this programm is made by seismosoft as well. seismostruct [12] uses the finite element method (fem) and can use material and geometrical nonlinearities of the structure. seismostruct [12] was used for calculations of the eigenvalues, eigenmodes and seismic response. • for the pushover analysis, a programm stab2dnl [13] was used. this programm is made by dr. uwe pfeiffer from university in hamburg. stab2d-nl [13] can compute 2d frames with the finite element method including material and geometrical nonlinearities. this software can consider the third-order-analysis as well. 3. dynamic properties of the structure 3.1. eigenmodes for the first rough estimation of eigenmodes, a simplified method which neglects all deformations of the cross-beam can be used. that means the crossbeam has an infinite stiffness. this simplification allows to analyse the structure as a system with two degrees of freedom. computational model is now a cantilever with two masses. the first eigenfrequency is f0,1 = 6.62 hz and the second eigenfrequency is f0,2 = 16.87 hz. eigenmodes were normalized to the roof level. the first eigenmode is defined as φ1 = (0.632;1)t and the second mode is φ2 = (-1.383;1)t. figure 5 shows the first and the second eigenmodes. figure 5. the first and the second eigenmode. 3.2. eigenmodes – fem (1.) frame with flexible beams (2.) frame with rigid beams for following calculations, an infinitely stiff model was chosen. this effect was defined with a very high 22 vol. 13/2017 comparison of methods used for seismic analysis of structures eigenmode 1 2 eigenfrequency [hz] 5.830 15.853 eigenperiod [s] 0.1715 0.0631 effective modal mass wxi/wx,tot [%] 92.50 7.49 σ 99.99 table 1. eigenvalue properties of the pliable stiff cross-beam. stiffness of the cross-beam, specifically with the enlargement of the e-modulus in factor 40. from this moment are eigenvalues almost constant. eigenmode 1 2 eigenfrequency [hz] 6.749 16.870 eigenperiod [s] 0.1482 0.0593 effective modal mass wxi/wx,tot [%] 95.16 4.83 σ 99.99 table 2. eigenvalue properties of the infinitely stiff cross-beam. 4. lateral force method of analysis this method is a static analysis using linear behaviour of the material, where horizontal forces represent the seismic load act to the joints of the structure. it is possible to use this method only for structures with small relevant eigenmodes. this method is used for a direct design of the cross-section, mostly from concrete and steel [14]. for this calculation, a proceeding according to [1], chapter 4.3.3.2 was used. the most important results are the design spectrum sd(t 1) = 2.4 m/s2 and the base shear force fb = 57.6 kn. these values approach to horizontal forces acting on all storeys. the force for the first storey is f1 = 24.16 kn and for the roof level is f2 = 33.44 kn. values φ1 and φ2 were assumed from the hand calculation in the chapter 3.1. final horizontal displacements are u1 = 0.00109 m for the first storey and u2 = 0.001724 m for the second one. 5. modal response spectrum analysis 5.1. hand calculation input values for the structure and the response spectrum are equivalent to calculations in chapters 2.2.2 and 3.1. in the first step, horizontal forces for the first eigenmode f1 = 22.98 kn, f2 = 31.78 kn and for the second eigenmode f1 = 4.32 kn, f2 = -2.73 kn were defined. in the second step, horizontal displacements for the first eigenmode u1 = 0.001037 m, u2 = 0.00003 m, and for the second eigenmode u1 = 0.00003 m, u2 = 0.000022 m were derived. from these results of the both eigenmodes, the total displacement was calculated. srss method was used for the combination of displacements and target results are u1 = 0.001038 m a u2 = 0.001639 m. results show that the effect of the second eigenmode is marginal. as a proof, a calculation of effective modal masses can be used – mef f1 = 22.815 and m ef f 2 = 1.184. 5.2. fem – linear horizontal displacements were analysed by the fem program with linear material behaviour of two versions: (1.) frame with flexible beams uk,max = 0.00209002 m, uk,min = -0.00230263 m. (2.) frame with rigid beams uk,max = 0.00157214 m, uk,min = -0.00158202 m. 6. time-history analysis the time-history analysis is another option, how the seismic response can be calculated. a model for this calculation is imposed by up to three different accelerograms which describe the behaviour for each direction. usually is this method used as non-linear analysis. to compare this method with two linear methods is necessary to use synthetic accelerograms. these accelerograms describe the response spectrum. ec8 [1] defines, that for the determination of the design seismic action, an average value from at least seven nonlinear calculations has to be used. by fever calculations, the worst result has to be chosen. calculation of motion equations is the main principle of this method and it is calculated by the modal analysis or by the direct integration. modal analysis allows the calculation with only linear material behaviour. with the direct integration is possible to use material non-linearities. 6.1. horizontal load this method is usually used for a calculation of any past known earthquake with defined accelerogram. for this comparison, accelerograms from the spectrum have to be defined. in fact, it is a reverse procedure. for the simplification, only one accelerogram was chosen. deformation calculated by this method has almost the same value as the result by the response spectrum method. the software seismoartif [11] was chosen to generate the optimal accelerogram. for the generation of synthetic accelerogram, a direct integration according to [15] (this method will be not described in this article) was used. 23 petr čada, jiří máca acta polytechnica ctu proceedings figure 6. acceleration. figure 7. velocity. figure 8. displacement. input values for the generation of synthetic accelerogram by programm seismoartif [11]: (1.) response spectrum according to chapter 2.2.2 (2.) earthquake parameters • inter-plate regime of far-field • moment magnitude – 5 • distance between the main event and the station – 5 km • soil type – generic soil (vs,30 = 310 km/s) (3.) generation of the synthetic accelerogram • smallest period of desired response spectrum – 0.02 s • largest period of desired response spectrum – 3 s • damping – 5 % • time-step – 0.01 s • pga – 0.0815g = 0.8 m/s 6.2. comparison to the response spectrum programm seismoartif [11] allows a comparison of time-history charts with the assigned response spectrum – figures 9, 10 and 11. 6.3. linear time-history analysis horizontal displacements were analysed by the fem program seismostruct[12] for two versions of the structure. (1.) frame with flexible beams uk,max = 0.00236767 m, uk,min = -0.00283672 m. (2.) frame with rigid beams uk,max = 0.001553 m, uk,min = -0.001545 m. 6.4. non-linear time-history analysis total horizontal displacement was calculated by the same software, but in this version, the non-linear material behaviour was used. (1.) frame with flexible beams uk,max = 0.00238527 m, uk,min = -0.00253367 m. 24 vol. 13/2017 comparison of methods used for seismic analysis of structures figure 9. comparison of the acceleration. figure 10. comparison of the velocity. figure 11. comparison of the displacement. (2.) frame with rigid beams uk,max = 0.00155074 m, uk,min = -0.00154217 m. 7. non-linear static (pushover) analysis pushover analysis is the last method of the ec8 [1] for the calculation of the seismic response. this nonlinear method is based on constantly acting self-weight and monotonically increasing horizontal load. in this method, plastic mechanism has to be defined by plastic regions. sufficient ductility has to be ensured by good construction design of plastic regions. remaining regions must have a big capacity and they have to stay elastic while plastic regions are overloaded. this procedure determines members for the energy dissipating. “soft” regions are for dissipating and “hard” are elastic [14]. 7.1. hand calculations of the frame with flexible beams by pushover analysis for determination of the pushover curve is necessary to monotonically increase the horizontal load up to the collapse of the structure or up to the maximal allowed horizontal displacement. according to [1] is necessary to applicate at least two different distributions of the lateral load (constant, linear, modal, . . . ). for a simple comparison, only one distribution of the lateral load was used. in this case a modal distribution of the first eigenmode was chosen. the pga is 0.8 m/s2. the product of the pga and the total weight defines the base shear force fb = 20.48 kn. the base shear force is distributed to the structure according to normalized eigenmodes. pushover curve describes the relation between the base shear force fb and the horizontal displacement d of the roof level. this capacity curve shows the yielding progress of the structure and regions of plastic hinges. for calculation of the monotonically increasing load, the software stab2d-nl [13] was chosen. iterative non-linear calculation determined the maximal load by collapse of the structure by 7.4-times of the assigned lateral load – f1 = 60.83 kn a f2 = 82.88 kn. lateral load was increased in 50 steps, respectively by 2.0 %. vertical load acts in all steps by 100 %. the collapse of the structure occurs by the depletion of the crosssection strength by the member nr. 3. this method determines the capacity curve for the member nr. 6 in the roof level. horizontal displacement of the multiple degree of freedom (mdof) system in the roof level before collapse is 37.7 mm. for the transformation to the single degree of freedom (sdof) system is necessary to 25 petr čada, jiří máca acta polytechnica ctu proceedings figure 12. example of the time-chart of the displacement. figure 13. capacity curve of the model with flexible beams. determine transformation factor γ. γ was counted according to procedure in [1], annex b, and is equal to 1.127. next step in this method is the determination of the deformation energy. deformation energy is the surface under the capacity curve of sdof system (figure 13). this surface was calculated by integration of sextic equation which was defined by programm excel (equation of the capacity curve). deformation energy em was transformed to the deformation energy of the sdof system by transformation factor γ. em* is equal to 2875.07 knmm. figure 14. determination of the idealized elastoperfectly plastic force-displacement relationship for the frame with flexible beams. determination of the idealized elasto-perfectly plastic force-displacement relationship is shown in figure 14. the period of the idealized sdof system was determined (t* = 0.362 s). with this period, the horizontal displacement can be calculated. because the period of the sdof is smaller than the tc-value of the response spectrum, a short period range can be considered. because the calculated displacement of the sdof was too far from the dm*, iterative solution was used. after the fifth iteration, the target displacement is equal to d∗et = 4.37 mm. the displacement was transformed by the transformation factor to the target displacement of the mdof system – dt = 4.92 mm. 26 vol. 13/2017 comparison of methods used for seismic analysis of structures figure 15. capacity curve of the model with rigid beams. 7.2. hand calculations of the frame with rigid beams by pushover analysis the weight of the system stays identical. distribution of lateral forces is changed with dependence on the stiffness. collapse of the stiff structure occurs by the same member by 7.3-times of the assigned lateral load – f1 = 59.641 kn a f2 = 81.76 kn. the capacity curve for the member nr. 6 in the roof level was determined (figure 15) horizontal displacement of the mdof system in the roof level before collapse is 31.4 mm. for the transformation to the sdof system is necessary to determine transformation factor γ. γ was counted according to procedure in [1], annex b, and is equal to 1.129. deformation energy em* was determined by the same procedure as in previous chapter. em* is equal to 2428.84 knmm. determination of the idealized elasto-perfectly plastic force-displacement relationship is shown in figure 16. the target displacement of the sdof system was determined as in previous chapter. six iterations were needed for the target displacement of the sdof – d*et = 3.02 mm. the transformed displacement of the mdof system is equal to dt = 3.41 mm. 8. comparison and conclusion all calculation methods for the seismic response according to ec8 [1] were compared. a simple model of the two-storeyed frame was used. the main line of the comparison was in the following way: • hand calculations of dynamical characteristics were made. figure 16. determination of the idealized elastoperfectly plastic force – displacement relationship for the frame with rigid beams. • these values were used for the hand calculation of the response from the spectrum. • results of the response spectrum were compared with software results. • an accelerogram for time-history analysis with identical displacements to the response spectrum method was determined • the chosen accelerogram was used for further calculations. • results were compared with lateral force method and pushover analysis. all results correspond with the theory that the stiffer structure approaches to smaller deformations. the lateral force method is the most conservative method for the seismic response. this method is used for simple symmetrical structures with predictable behaviour. the lateral force method is the simplest method and it is not necessary to use special software or complicated calculations. a disadvantage of this method is only linear material behaviour. the next linear method is the response spectrum method. this method gave us better results. this is a good choice for more complicated structures. members of the structure are with this analysis not so much oversized. it is widely used in the practice. the non-linear behaviour changes results significantly. non-linear methods according to ec8 [1] are the time-history analysis and the pushover analysis. both of them are particular and it is necessary to use of some specialized software. the time-history analysis is suitable for the analysis of past known earthquakes. the pushover analysis is better for design of new structures or reconstructions. all results are compared in table 3. identical results of the horizontal displacement between linear calculations and dynamic calculation were achieved. on the other hand, a bigger difference between the pushover analysis and other methods was found out. this difference implies to a requirement to be cautious by this method and to focus on initial 27 petr čada, jiří máca acta polytechnica ctu proceedings calculation method horizontal displacement of the structure [mm] cross-beam stiffness linear non-linear lateral force method flexible rigid 1.724 d response spectrum method flexible 2.303 a rigid 1.582 a, 1.639 d time-history analysis flexible 2.837 a 2.534 b rigid 1.545 a 1.543 b pushover analysis flexible 4.92 b,c rigid 3.41 b,c (a) calculation by software seismostruct [12] (b) calculation by software stab2d-nl [13] (c) hand calculation table 3. comparison of calculations of the seismic response. values. for these reasons, it is necessary to focus more on this method and to study another new conditions and inputs of its application. acknowledgements the authors gratefully acknowledge support from the czech technical university in prague, project sgs17/043/ohk1/1t/11 development and application of advanced algorithms for numerical analysis and modelling of structures and materials. references [1] en 1998-1: eurocode 8: design of structures for earthquake resistance: part 1: general rules, seismic actions and rules for buildings, 2004. [2] p. bisch, e. carvalho, h. degee, et al. eurocode 8: seismic design of buildings: worked examples. 1 lisbon: publications office of the european union 2012. [3] a. faroughi, m. hosseini. simplification of earthquake accelerograms for quick time history analyses by using their modified inverse fourier transforms. procedia engineering 14:2872–2877, 2011. doi:10.1016/j.proeng.2011.07.361. [4] e. kalkan, a. k. chopra. practical guidelines to select and scale earthquake records for nonlinear response history analysis of structures. us geological survey open-file report 2010. [5] p. fajfar, m. fischinger. n2 a method for non-linear seisimic analysis of regular buildings. proceedings of ninth world conference on earthquake engineering pp. v111–v116, 1988. doi:10.1002/eqe.144. [6] a. k. chopra, r. k. goel. a modal pushover analysis procedure for estimating seismic demands for buildings. earthquake engineering 31(3):561–582, 2002. doi:10.1002/eqe.144. [7] p. fajfar. a nonlinear analysis method for performance based seismic design. earthquake spectra 16(3):573–592, 2000. [8] v. kilar, p. fajfar. simplified push-over analysis of building structures. eleventh world conference on earthquake engineering acapulco (1011):2872–2877, 1996. [9] m. kreslin, p. fajfar. the extended n2 method taking into account higher mode effects in elevation. earthquake engineering 40(14):1571–1589, 2011. doi:10.1002/eqe.1104. [10] en 1992-1-1: eurocode 2: design of concrete structures: part 1-1: general rules and rules for buildings, 2004. [11] programm seismoartif. http://www.seismosoft.com/seismoartif. [12] programm seismostruct. http://www.seismosoft.com/seismostruct. [13] programm stab2d-nl. http://www.u-pfeiffer.de/. [14] h. bachmann. erdbebensicherung von bauwerken. basel boston berlin: birkhäuser verlag, 2002. [15] b. halldorsson. calibration of the specific barrier model to earthquakes of different tectonic regions. bulletin of the seismological society of america 95(4):1276–1300, 2005. doi:10.1785/0120040157. 28 http://dx.doi.org/10.1016/j.proeng.2011.07.361 http://dx.doi.org/10.1002/eqe.144 http://dx.doi.org/10.1002/eqe.144 http://dx.doi.org/10.1002/eqe.1104 http://www.seismosoft.com/seismoartif http://www.seismosoft.com/seismostruct http://www.u-pfeiffer.de/ http://dx.doi.org/10.1785/0120040157 acta polytechnica ctu proceedings 13:21–29, 2017 1 introduction 2 structure and load 2.1 model of the structure and material parameters 2.2 load 2.2.1 vertical load 2.2.2 seismic load 2.3 software 3 dynamic properties of the structure 3.1 eigenmodes 3.2 eigenmodes – fem 4 lateral force method of analysis 5 modal response spectrum analysis 5.1 hand calculation 5.2 fem – linear 6 time-history analysis 6.1 horizontal load 6.2 comparison to the response spectrum 6.3 linear time-history analysis 6.4 non-linear time-history analysis 7 non-linear static (pushover) analysis 7.1 hand calculations of the frame with flexible beams by pushover analysis 7.2 hand calculations of the frame with rigid beams by pushover analysis 8 comparison and conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0074 acta polytechnica ctu proceedings 15:74–80, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app the influence of boundary conditions on the response of underground structures subjected to earthquake veronika pavelcová, tereza poklopová, tomáš janda, michal šejnoha∗ czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: sejnom@fsv.cvut.cz abstract. the paper deals with the prediction of the response of a real underground structure subjected to earthquake. a fully dynamic analysis is carried out in the geo5 fem program using the finite element method. limiting our attention to a two-dimensional analysis we focus on the implementation of special boundary conditions along the vertical edges of the computational model. a simple study is carried out first to show that incorrectly applied boundary conditions may significantly influence the actual design of underground structures loaded by vertically propagating shear waves. this study promotes the combination of so called free-field and static boundary conditions as demonstrated on a simple example. keywords: earthquake, shear weave, accelerogram, finite element method, free-field boundary, absorbing boundary, fixed boundary. 1. introduction crowded cities, pushing the need for fast and safe transportation, are the main cause of ever faster growing number of underground structures. such structures, built in densely populated areas, often experience a high degree of complexity. to maintain reliable performance of tunnels during their life time calls for reliable designs. these are typically based on detailed finite element simulations. in this regard, the preliminary design typically adopts a certain simplification such as a dimensional reduction from 3d (three-dimensional) to 2d (two-dimensional). while simplified 2d formulations attempting to account for out-of-plane deformations do exit [1–3], the majority of practical engineers rely on a fully 2d plane-strain analysis employing the convergence confinement method (2dcc) either in deterministic [4] or stochastic [5] environment. owing to its popularity and availability in various geotechnical softwares, e.g. geo5 tunnel [6], the 2dcc analysis is often the choice even if addressing the tbm (tunnel boring machine) based excavation strategies. while in most applications the static loading conditions play the principal role in the design of tunnel lining or in assessing a potential failure during tunnel construction and when in service, the dynamic analysis in regions prone to the occurrence of earthquake may prove important. for a general and clear introduction to this subject the reader is referred to an excellent monograph by i. towhata [7]. when limiting our attention to the determination of lining forces caused by earthquake it is often sufficient to carried out a simplified pseudo-static analysis with the help of either analytical [8, 9] or fem based methods [10–12]. the principal concern in connection to a fully dynamic analysis, regardless whether 2d or 3d, is the representation of infinite boundary in the horizontal direction. assuring that horizontally propagating waves are not reflected back to the domain requires implementation of appropriate boundary conditions along vertical boundaries. these are typically in the form of viscous dampers relating the boundary tractions to velocities of the propagating waves. clearly, absorbing these waves at the boundary has the same effect as if transmitting them with no restrictions. the most widely used absorbing/transmitting boundary conditions are those introduced by lysmer and kuhlemeyer [13]. if represented by these boundary conditions the boundary absorbs the entire wave propagating across this boundary. when modeling the response of structure to earthquake it appears more appropriate to absorb only the portion of energy associated with the perturbation from so called free field conditions, the state preceding any construction. such perturbation can be attributed to, e.g. the built above-ground or underground structure such as embankment or tunnel. such boundary conditions, termed henceforth the free-field boundary conditions, were put forward by zinkiewicz, et al. [14]. to acquire the necessary free field response they build upon the solution of one-dimensional (1d) free field column problem [10, 12, 14]. point out that free-field boundary conditions should become active only if there are other than vertically propagating waves. thus to properly represent the results of 1d free field column analysis within a 2d environment calls for other types of boundary conditions that assure, in case of verti74 http://dx.doi.org/10.14311/app.2018.15.0074 http://ojs.cvut.cz/ojs/index.php/app vol. 15/2018 influence of boundary conditions in earthquake analysis cally propagating shear weaves, the stress state the same along every vertical section in the 2d model as derived from 1d analysis. starting with a brief introduction to a theoretical background in section 2 we continue in section 3 with a simple parametric study on the effect of various types of potential boundary conditions applied along the lateral boundaries of the analyzed domain. the most appropriate boundary conditions are then tested on an illustrative example of ground excavation in section 4. the essential findings are finally summarized in section 5. 2. theoretical background suppose that a simple homogeneous rectangular domain is loaded at its boundary by a vertically propagating shear weave as schematically shown in fig. 1(a). the terms uu,ud stand for the upward and downward traveling waves, respectively. we further assume an infinitely stiff base rock so the total motion in the form of outcrop motion can be applied at the bottom boundary. (a) (b) figure 1. (a) schematic representation of loading through outcrop motion, (b) loading and boundary conditions applied to 1d and 2d domains. the displacement field in space and time can then be decomposed as u(x, t) = u0(t) + ur(x, t), (1) where at the bottom boundary it holds u(x, t)|y=0 = u0(t), ur(x, t)|y=0 = 0. (2) the discretized form of the equation of motion then reads mür(t) + cu̇r(t) + kur(t) = −mü0(t), (3) where m, c, k are the mass, damping and stiffness matrices, respectively. note that no material damping is assumed so that the damping matrix c collects only the viscous dampers applied along the domain lateral boundaries as shown in fig. 1(b). therein, the viscosity cp = √ eoed/ρ corresponds to the pressure wave velocity whereas cs = √ g/ρ is associated with the shear wave velocity; ρ,eoed,g stand for the soil density and oedometric and shear moduli, respectively. further details can be found in [13–15]. the time integration of eq. (3) is typically performed with the help of newmark integration scheme. as already mentioned in the introductory part the analysis of 1d free field column problem seen in fig. 1(b) is needed to provide free-field velocities u̇f fr and shear tractions py = τxy. in the present study, a simple impulse plotted in fig. 2 was considered in both 1d and 2d analyses. (a) (b) figure 2. prescribed impulse: (a) acceleration, (b) displacement. to eliminate any potential errors caused by finite element discretization we assumed an optimal element size defined such that for the selected time step of 0.01s the shear wave passes through only just one element, i.e. lelem = cs∆t = √ g/ρ∆t =√ 28.6 × 106/1960 · 0.01 = 1.2 m. the potential error arising from discretization if deviating from an optimal element length is clearly visible in fig. 3. thus only the optimal element length was adopted when performing all relevant calculations. one such analysis considers a 1d free-field column to provide results to be reproduced by a 2d analysis discussed in the next section. example of time variation of horizontal displacements of two selected points (on surface and at the center of the domain) is plotted in fig. 4 for illustration. 75 v. pavelcová, t. poklopová, t. janda, m. šejnoha acta polytechnica ctu proceedings (a) (b) figure 3. influence of finite element discretization. distribution of horizontal displacement u along vertical direction at time: (a) t = 0.4 s, (b) t = 4.0 s. (a) (b) figure 4. time variation of horizontal and vertical displacements u,v: (a) on surface, (b) at the center of the domain. as seen, the vertical displacements are clearly zero as only the horizontal displacements in terms of acceleration ü0(t) are prescribed to the entire domain. owing to the fixed boundary ur(0, t) = 0 and no material damping, the wave is free to bounce back to the domain when arriving at the bottom boundary. also note that the top boundary, free of shear stress, amplifies the propagating shear wave at the boundary two times in comparison to the wave within the domain, see e.g. [7]. 3. parametric study on the influence of lateral boundary conditions to see how well may various boundary conditions represent the free-field response we perform a simple parametric study. to that end, a rectangular homogeneous domain with the shear modulus g = 28.6 mpa, the poisson ratio ν = 0.4 and the density ρ = 1960 kg/m3 is considered. as in the previous section we reserve ourselves to fixed boundary at the bottom and loading conditions specified in fig. 2. in particular, four types of boundary conditions plotted in fig. 5 are examined in a sequel. (a) (b) (c) (d) figure 5. selected boundary conditions along lateral edges: (a) fixed vertical displacement v(t) = 0 (bc1), (b) free boundary-no restrictions (bc2), c) freefield boundary conditions (bc3), d) static boundary conditions py (t) prescribed (bc4). 3.1. fixed boundary conditions (bc1) with reference to the results provided by free field column analysis we begin by prescribing a kinematic constrain in the vertical direction, fig. 5(a). such boundary conditions assure zero vertical displacement along lateral edges while generating, through the reaction forces, the expected non-zero shear stress τxy. as evident from fig. 6, these are the correct boundary conditions producing the same response as the 1d free field column analysis. this is supported by the distribution of the shear stress found at the end of analysis in fig. 10(a). unfortunately, such boundary conditions are generally not acceptable as they do not allow for transmitting the horizontally propagating wave caused, e.g. by some structural perturbation. thus application of such boundary conditions could only be justify in cases when the lateral boundaries are sufficiently far 76 vol. 15/2018 influence of boundary conditions in earthquake analysis from the point of interest, e.g. tunnel, and the reflected horizontal wave would not interfere with the incoming wave. however, such a requirement may lead to computationally demanding geometrical models. creating unstructured meshes with coarsening in the vicinity of lateral boundaries as typical of static analysis may not be the solution bearing in mind the mutual connection between the integration time step and the element size. such a boundary condition is, therefore, generally not recommended. (a) (b) figure 6. time variation of horizontal and vertical displacements u,v: (a) on surface, (b) at the center of the domain. 3.2. free boundary (bc2) just for illustration we consider next a free boundary with no restrictions. similar to terrain surface, such a boundary requires the shear stresses along lateral edges be zero. this in turn generates unacceptable vertical displacements within the whole domain and yield a nonuniform distribution of the shear stress as evident in fig. 10(b). the evolution of ineligible vertical displacements is further seen in fig. 7(b), which generates a strong deviation of the horizontal displacement in every section of the 2d domain when compared to the 1d free field column analysis. for the sake of clarity, the influence of free boundary on the results pertinent to surface point on the left lateral edge are reported in fig. 7(a). the free boundary is therefore a totally incorrect boundary condition, even if the lateral boundaries are sufficiently far. (a) (b) figure 7. time variation of surface displacements: (a) horizontal displacement u, (b) vertical displacement v at the edge. 3.3. free-field boundary conditions (bc3) recall that free-field boundary conditions in fig. 5(c) should allow for transmitting the structural or material perturbation only causing, as will be seen in section 4, e.g. a horizontally propagating shear weave. however, no such perturbation is present in the present example. so both viscous dampers (the horizontal one transmitting the horizontally propagating pressure wave, the vertical one transmitting the horizontally propagating shear wave) should remain inactive during the analysis. however, the formulation of damper does not distinguish between the type and direction of the propagating wave. with reference to the previous section it is therefore expected the vertical damper to become active as attempting to damp the difference between the 1d (v = 0) and 2d (v 6= 0) analyses. this is evident from fig. 8(b) showing the variation of surface vertical displacement being 10 times smaller than in case of free boundary examined in section 3.2. the results, fully compatible with the 1d free field column analysis, could thus be achieved by setting the shear viscosity cs to infinity, see [15]. clearly, the infinite viscosity acts in this case as a penalty forcing the velocity of vertical displacement to zero thus in turn generating, similar to section 3.1, the nonzero shear stresses. nevertheless, even with damper of finite viscosity the horizontal displacements in fig 8(a) are repre77 v. pavelcová, t. poklopová, t. janda, m. šejnoha acta polytechnica ctu proceedings sented fairly well. similarly, the distribution of shear stresses displayed in fig. 10(c) is considerably improved. but keep in mind that the action of vertical damper is in this case purely artificial, as it is activated for purpose, for which it is not actually designed. (a) (b) figure 8. time variation of surface displacements: (a) horizontal displacement u, (b) vertical displacement v at the edge. 3.4. static boundary conditions (bc4) intuitively, and recalling the discussion in section 3.1, the most appropriate boundary conditions to reproduce the 1d free field column analysis are static boundary conditions represented by the prescribed shear tractions as shown in fig. 5(d). it has been demonstrated in [15] that such a boundary conditions yield identical displacements and stresses as produced with the bc1 type of boundary conditions, even if combined with the free-field boundary conditions (bc3), which in this case remain inactive. figure 9. periodic boundary conditions. it is interesting to point out that similar results would be derived with the application of periodic boundary conditions [14] schematically depicted in fig. 9. these boundary conditions, widely used in the field of homogenization [16], are essentially the tying constrains which force the homologous displacements on the opposite sides of the analyzed domain the same. [kpa] (a) [kpa] (b) [kpa] (c) figure 10. distribution of shear stress τxy at the end of analysis for different lateral boundary conditions: (a) bc1 and bc4, (b) bc2, (c) bc3. 4. example to illustrate the effect of combined free-field and static boundary conditions to demonstrate the applicability of the proposed boundary conditions we consider a geometrical model plotted in fig. 11(a) and loaded by the same impulse as used in the previous section, recall fig. 2. in particular, the effect of perturbation caused by the 90◦ wedge will be examined. in this particular example it is expected that the upward traveling shear wave, when hitting the wedge boundary, will be reflected in the 90◦ angle to become a horizontally propagating shear wave as illustrated in figs. 11(b,c). to appreciate the action of free-field boundary conditions we begin with the static boundary conditions only and follow the time evolution of the vertical and horizontal components of the displacement vector linked to one particular point marked in fig. 11(a) 78 vol. 15/2018 influence of boundary conditions in earthquake analysis (coordinates [40,30]). the results appear in fig. 12. it is clear from fig. 12(b) that in view of the horizontally propagating shear wave the horizontal boundary behaves as a free boundary and reflects the wave back into the domain. because the bottom boundary is assumed fixed, the wave is trapped in the domain, and in case of no material damping will continue to infinitely move between the two boundaries. the fact that the impulse is applied along the entire bottom boundary promotes interaction of individual waves which in turn is manifested by a continuous rise of the wave amplitude. (a) (b) (c) figure 11. testing boundary conditions: (a) geometrical model, (b) static boundary conditions assumed only, (c) combination of free-field and static boundary conditions. a considerably different response is observed when properly combining the static boundary condition with the free-field boundary condition, see the red curves in figs. 12. as evident from fig. 12(b), the latter boundary condition is able to transmit, damp, the horizontally propagating incoming wave across the lateral edges, so it is not reflected back. the subsequent vibration is again the result of interaction of the neighboring waves. the continuous vibration of the tracked point in the horizontal direction, see fig. 12(b), is caused by continuously traveling shear wave in the vertical direction caused again by the presence of fixed boundary assumed at the bottom of the model. to bring the results closer to reality one would have to either introduce the material damping or the absorbing boundary condition at the bottom boundary, see e.g. [11, 12, 15]. 5. conclusion application of various boundary conditions along vertical edges of a rectangular domain was critically exam(a) (b) figure 12. time variation of the displacement of the marked point: (a) horizontal displacement u, (b) vertical displacement v. ined in this paper. it was proved first in section 3 that static boundary conditions are the correct boundary conditions to represent the 1d free field column analysis in 2d simulations. these boundary conditions were then combined in section 4 with the free-field boundary conditions to allow for properly transmitting the horizontally propagating shear weave across the lateral boundaries perturbed in this particular case by an excavated wedge. the ability of free-field boundary conditions to absorb such waves has been confirmed. however, the performance of these absorbing conditions in cases of incident angles other than 90◦ is still an open issue and will be examined in our future studies. acknowledgements the support provided by the sgs project no. sgs18/037/ohk1/1t/11 and the tačr project no. te01020168 is gratefully acknowledged. references [1] t. janda, m. šejnoha, j. šejnoha. application of converegence measurements in 2d analysis of tunnel excavation in comparison with convergenece confinement method. tunel 19(4):52–57, 2010. [2] t. janda, m. šejnoha, j. šejnoha. modeling successive excavation within two dimensional finite element mesh. acta geodynamica et geomaterialia 8(1):69–78, 2011. [3] t. janda, m. šejnoha, j. šejnoha. modeling of soil structure interaction during tunnel excavation: an engineering approach. advances in engineering 79 v. pavelcová, t. poklopová, t. janda, m. šejnoha acta polytechnica ctu proceedings software 62-63:51–60, 2012. special issue dedicated to professor zdeněk bittnar on the occasion of his seventieth birthday: part i. [4] t. svoboda, d. mašín. comparison of displacement field predicted by 2d and 3d finite element modelling of shallow natm tunnels in clays. geotechnik 34(2):115–126, 2011. [5] t. janda, m. šejnoha, j. šejnoha. applying bayesian approach to predict deformations during tunnel construction. international journal for numerical and analytical methods in geomechanics 42:1765–1784, 2018. [6] fine ltd. geo 5 civil engineering sowftware. [online]: www.fine.cz. [7] i. towhata. geotechnical earthquake engineering. springer-verlag berlin heidelberg, 2008. [8] j. penzien. seismically induced racking of tunnel linings. earthquake engineering and structural dynamics 29:683–691, 2000. [9] y. hashash, j. hook, b. schmidt, j. yao. seismic design and analysis of underground structures. tunneling and undergrounding space technology 16:247–293, 2001. [10] d. kučera. posouzení geotechnické konstrukce na seismické zatížení metodou konečných prvku. master’s thesis, czech technical university in prague, faculty of civil engineering, 2017. [11] t. poklopová. evaluation of real underground structure subjected to earthquake pseudostatic analysis. bachalor’s thesis, czech technical university in prague, faculty of civil engineering, 2018. in czech. [12] t. poklopová, v. pavelcová, t. janda, m. šejnoha. evaluation of real underground structure subjected to earthquake pseudostatic approach. acta polytechnica 2018. accepted. [13] j. lysmer, r. kuhlemeyer. finite dynamic model for infinite media. journal of the engineering mechanics division, proceedings of the american society of civil engineers 95(em4):859–877, 1969. [14] o. zinkiewicz, n. bicanic, f. shen. generalized smith boundary a transmitting boundary for dynamic computations, vol. 207. innme swansea, 1986. [15] v. pavelcová. evaluation of real underground structure subjected to earthquake fully dynamic analysis. bachalor’s thesis, czech technical university in prague, faculty of civil engineering, 2018. in czech. [16] m. šejnoha, j. zeman. micromechanics in practice. wit press, southampton, boston, 2013. 80 acta polytechnica ctu proceedings 15:74–80, 2018 1 introduction 2 theoretical background 3 parametric study on the influence of lateral boundary conditions 3.1 fixed boundary conditions (bc1) 3.2 free boundary (bc2) 3.3 free-field boundary conditions (bc3) 3.4 static boundary conditions (bc4) 4 example to illustrate the effect of combined free-field and static boundary conditions 5 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0148 acta polytechnica ctu proceedings 13:148–151, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app modal analysis for revision of a fem model of a steel truss beam michal venglar∗, milan sokol, monika marfoldi slovak university of technology in bratislava, faculty of civil engineering, department of structural mechanics, radlinskeho 11, 810 05 bratislava, slovak republic ∗ corresponding author: michal.venglar@stuba.sk abstract. the paper deals with a preparation of a complex fem model for a local damage detection. the initial verified and validated three-dimensional fem model of a steel truss bridge in laboratory is revised step-by-step to achieve the accurate model according to the experimental model. the emphasis is on modelling of the joints with 4 rivets and modelling of correct boundary conditions, as well as mass parameters and cross-section dimensions.a modal analysis of the structure is performed in fem software. many experimental measurements were made to correctly revise the fem model. the calculated natural frequencies are compared with the measured ones. in addition, mode-shapes from the calculation are validated with the measured mode-shapes. the difference between the prepared fem model and the measured specimen is small enough after a few steps of tuning. the verified, validated and revised numerical model can be used in future for a local damage detection. keywords: modal analysis, revision of fem model, experimental measurements, rivet joint connection. 1. introduction civil infrastructure like bridge structures deteriorate with time due to various reasons. the reasons are, e.g., fatigue failure, extreme events, etc. in addition, many of these structures worldwide are currently nearing the end of their proposed design life. this situation can result into a necessity of major rehabilitation. according to the 2017 infrastructure report card [1] produced by the american society of civil engineers (asce) almost 4 in 10 bridges are over 50 years and older. this means that bridge structures in the usa are 43 years old on average. situation in slovakia is almost the same. because of mentioned state, structural health monitoring (shm) and system identification of bridges is currently reaching popularity among research teams around the world [2–7]. moreover, increasing safety demands of new-constructed bridges and buildings confirms the interest. the shm of bridge structures can help to prevent from the stated situations. therefore, this paper deals with one of the first parts of shm (preparation of a fem model of bridge specimen). 2. laboratory specimen a steel truss bridge was scaled to approximately 1:15. the whole length of the specimen was 3.4 m. the cross-section of the truss is closed (through truss) with a width of 0.435 m and a height of about 0.8 m. the cross-section of upper and lower chords has been chosen as a square hollow section shs 15 × 1.5 mm. diagonals of the truss beam form a 63 °angle with the bottom and/or with the upper chords. the diagonals at the joints have the same cross-section as the chords (shs 15 × 1.5 mm) that changes to a full round bar figure 1. the boundary conditions. with a radius of 12 mm. floor beams consists of two shs 15×1.5 mm bars which have been welded together. rods with a diameter of 6 mm are used for top and bottom lateral bracing. steel components included the following nominal properties: elastic modulus 200 gpa and poisson ratio equals to 0.3. the experimental model has also been weighed for comparison with the numerical model. 3. experimental measurements the used boundary conditions of the experimental model were as a simply supported beam. the truss beam was supported at 4 places. you can see it in figure 1. accelerometers mmf ks901.100 were placed at 13 points. magnetic mounting was used because of pos148 http://dx.doi.org/10.14311/app.2017.13.0148 http://ojs.cvut.cz/ojs/index.php/app vol. 13/2017 modal analysis for revision of a fem model... figure 2. the natural frequencies from initial measurements. figure 3. the numerical model. sibility to change place of sensor [8] in accordance with investigated mode-shape. the air temperature reached approximately 20 °c during the measurements and it was measured because of possible influence on repeated dynamic measurements in the future. the first three natural frequencies from measurements are shown in figure 2. analysis of the measurement data was done in software national instruments labview and modalview r2. 4. model revision in accordance with [9], the comparison of experimental outcomes and simulation outcomes was done. the simulation outcomes were prepared in the fem code where the numerical model (figure 3) was made from shell, beam elements and mass elements was also used for lumped mass. firstly, the model was verified and validated. then, the experimental outcomes were acquired in accordance with section 3. as it was mentioned in section 2, the weight of the laboratory specimen with accessories was compared figure 4. revised boundary conditions. with the numerical model. the numerical model has also considered the weight of the used accelerometers mmf ks901.100. at the beginning, the whole mass did not correspond to the mass of numerical model. then, cross-sectional dimensions and mass of individual elements of the system were carefully measured. the occurred difficulty during the process of validation was different thickness of individual cross-sections. finally, the whole mass was adjusted to weight of the experimental specimen. another difficulty which occurred during the validation and following revision of the model represents the boundary conditions. after many simulations, the boundary conditions were achieved by modelling l-profiles (figure 4) with the same length placed in four corners of the specimen. modelling of the hinge joints by using rivets (figure 5) was done as the final step. the rivets were modelled using a single-axis element that allows to specify 149 m. venglar, m. sokol, m. marfoldi acta polytechnica ctu proceedings figure 5. rivets in the numerical model. figure 6. the first calculated mode-shape: view (top) and cross-section (bottom). the outer diameter and thickness of the cross-sectional area. the thickness of the used rivets influenced each of the global natural frequency as well as frequencies of local vibrations of diagonal members. the most suitable value of the thickness was equal to 0.45 mm. 5. results of revision after verification, validation and revision of the numerical model, final modal analysis was done. the first natural frequency represents 9.37 hz. when the structure is deformed in its the first mode-shape, the entire upper part of the structure with diagonals moved in the direction of the y axis, as you can see in figure 6. for a comparison with the first calculated modeshape, the modal analysis was done in modalview r2 software. the better results was achieved in the figure 7. the first measured mode-shape: view (top) and cross-section (bottom). fem code, because modalview r2 software used many times less elements. but generally, the measured mode-shape (figure 7) is comparable to the calculated one. the measured frequency was approximately 9.35 hz. the frequency is also in good conformity. the comparison of another natural frequencies from the measurements and the numerical analysis is in figure 8. the first two mode-shapes and the fourth represent global mode-shapes. the third mode-shape is represented as vibration of diagonal members, so it is local mode-shape. the fifth and higher mode-shapes are also local modes. 150 vol. 13/2017 modal analysis for revision of a fem model... figure 8. comparison of measured and calculated frequencies. 6. summary and conclusions the laboratory investigation was utilized to test the numerical model. after mentioned revisions in the numerical model, the obtained results were satisfactory. the compared mass was equalled. the measured and calculated natural frequencies showed only small inaccuracies. the mode-shapes from the numerical model and the measurements gave us comparable outcomes. the tuned numerical model can be used for another calculations and measurements to detect local damages in joints. acknowledgements this paper was supported by the slovak research and development agency (srda), i.e. a grant from research program no. apvv-0236-12; it was also created with the support of the ministry of education, science, research and sport of the slovak republic within the research and development operational program for the project university science park of stu bratislava, itms 26240220084. it was also supported by grant from research program young researcher 2017. references [1] american society of civil engineers. the 2017 infrastructure report card. https://www.infrastructurereportcard.org/ wp-content/uploads/2016/10/ 2017-infrastructure-report-card.pdf. [2] z. orban. increasing the reliability of the assessment iof masonry archbridges by non-destructive testing. pollack periodica 1(3):45–56, 2006. doi:10.1556/pollack.1.2006.3.4. [3] s. c. siriwardane. vibration measurement-based simple technique for damage detection of truss bridges: a case study. case stud eng fail anal 4:50–58, 2015. doi:10.1016/j.csefa.2015.08.001. [4] k.-c. chang, c.-w. kim. modal-parameter identification and vibration-based damage detection of a damaged steel truss bridge. eng struct 122:156–173, 2016. doi:10.1016/j.engstruct.2016.04.057. [5] j. li, h. hao. health monitoring of joint conditions in steel truss bridges with relative displacement sensors. measurement 88:360–371, 2016. doi:10.1016/j.measurement.2015.12.009. [6] a. bayrakattar, a. c. altunisik, t. turker. structural health assessment and restoration procedure of an old riveted steel arch bridge. soil dyn earthquake eng 83:148–161, 2016. doi:10.1016/j.measurement.2015.12.009. [7] m. sokol, m. venglar. increasing the reliability of the assessment iof masonry archbridges by non-destructive testing. pollack periodica 12(3), 2017. [in press]. [8] j. s. wilson. test and measurement. first printing. newnes/elsevier, burlington, mass, 2009. [9] b. h. thacker, s. w. doebling, f. m. hemez, et al. concepts of model verification and validation. technical report, 2004. [la-14167-ms], https: //www.osti.gov/scitech/servlets/purl/835920. 151 https://www.infrastructurereportcard.org/wp-content/uploads/2016/10/2017-infrastructure-report-card.pdf https://www.infrastructurereportcard.org/wp-content/uploads/2016/10/2017-infrastructure-report-card.pdf https://www.infrastructurereportcard.org/wp-content/uploads/2016/10/2017-infrastructure-report-card.pdf http://dx.doi.org/10.1556/pollack.1.2006.3.4 http://dx.doi.org/10.1016/j.csefa.2015.08.001 http://dx.doi.org/10.1016/j.engstruct.2016.04.057 http://dx.doi.org/10.1016/j.measurement.2015.12.009 http://dx.doi.org/10.1016/j.measurement.2015.12.009 https://www.osti.gov/scitech/servlets/purl/835920 https://www.osti.gov/scitech/servlets/purl/835920 acta polytechnica ctu proceedings 13:149–152, 2017 1 introduction 2 laboratory specimen 3 experimental measurements 4 model revision 5 results of revision 6 summary and conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2016.5.0069 acta polytechnica ctu proceedings 5:69–73, 2016 © czech technical university in prague, 2016 available online at http://ojs.cvut.cz/ojs/index.php/app deployment of ertms in czech republic jiří vopava∗, mária jánešová, radek kratochvíl ctu in prague, czech technical university in prague – faculty of transportation sciences, konviktská 20, 110 00, prague ∗ corresponding author: jiri.vopava@gmail.com abstract. the article describes current situation and expected deployment of international standard for automatic rail vehicle protection in czech republic for period 2014 – 2020. article introduces expected status in 2020 and subsequent implementation of commitments until 2026. it introduces expected result of gradual implementation for tracks and vehicles. it also displays brief deployment overview of european rail traffic management system in european union countries for each level of european train control system. keywords: ertms, etcs, gsm-r, interoperability. 1. introduction the ertms is automatic train protection which represents basic part of interoperability in european rail system. the term is abbreviation of "european rail traffic management system" and it was defined to synchronize nationals systems to the one common compatible system. urgency of the unification and cooperation between european countries caused an increase in freight which had negative impact for environment and ecology. the implementation of interoperability and automatic train protection across european countries allows competition to air and car freight. implementation of the standard is business initiative that transforms the operation of the tracks, enhance safety, capacity, performance and reliability. it also have impact on reduce of cost of operation and maintenance. ertms is divided into two basic components. the first one is etcs "european train control system" and second gsm-r "global system for mobile communication for railway". the ertms allow one compatible system for crossborder traffic what reduce cost for implementation several regional stand-alone systems. the synchronization of automatic train protection increase competitiveness of the european rail sector. gsm-r "global system for mobile communications – railway" is a radio system based on gsm standard for providing of voice and data communication between vehicle and track. etcs "european train control system", is an automatic train protection system to replace the existing national systems. the etcs replaces about twenty different automatic train protections across european track. it can be configured to operate in one of the following application levels to enable etcs equipped vehicles to operate in not etcs equipped system where the safe movement of the vehicle is controlled by the default national control system: figure 1. etcs l1 [1] • level 0 – etcs equipped vehicles operating on lines without etcs system, • level 1 – an conventional line-side signal is passed to the train via a switched transmitter on the track, repeating the indication from the conventional signalling system, which continues to be used to support safe train separation. it is the figure 1, • level 2 – movement authority is passed by the gsm-r radio network from a radio block centre to the train. conventional train detection systems are utilized in conjunction with interlocking systems used to enforce safe train separation. it is the figure 2, • level 3 – it builds on level 2, but increases safe train separation using critical data from the train. it is the figure 3. 2. deployment of ertms in czech republic the czech republic contributing to the fulfillment of legal solutions arising from following documents: • directive 2008/57/ec of the european parliament and of the council 69 http://dx.doi.org/10.14311/app.2016.5.0069 http://ojs.cvut.cz/ojs/index.php/app j. vopava, m. janesova, r. kratochvil acta polytechnica ctu proceedings figure 2. etcs l2 [1] figure 3. etcs l3 [1] of 17 june 2008 on the interoperability of the rail system within the community • decisions 2012/88/eu, commission decision of 25 january 2012 on the technical specification for interoperability relating to the control-command and signalling subsystems of the trans-european rail system (2012/696/eu a 2015/14/eu) • regulation (eu) no 1315/2013 of the european parliament and of the council of 11 december 2013, on union guidelines for the development of the transeuropean transport network and repealing decision no 661/2010/eu • regulation (eu) no 1316/2013 of the european parliament and of the council of 11 december 2013, establishing the connecting europe facility, amending regulation (eu) no 913/2010 and repealing regulations (ec) no 680/2007 and (ec) no 67/2010 the deployment is defined in document "národní implementační plán ertms" released by ministry of transport. the document defines maximum amount of payloads for each project gsm-r and etcs l2 with respect to co-financing from ertms funds. it also defines expected deployment schedule of each project. between 2005 and 2014 was equipped 1132 km of tracks across of the czech republic. it is connected to figure 4. map of gsm-r implementation [2] figure 5. map of etcs implementation [2] the gsm-r sždc, db, öbb, prorail and the home network operator. there are now mobile terminals equipped with gsm-r all čd vehicles used in regular operation. almost all čd vehicles used in regular operation are now equipped with gsm-r terminals. the first stage with fully deployed etcs system at 2021 will be track "praha – česká třebová – brno – břeclav". following tables 1, 2, 3 display figures for expected schedule of deployment of gsm-r, etcs l2 and vehicle on-board equipment for period 2014 – 2020 (– 2026). the figure 6 below visualizes data from tables above. it displays total amount of deployments for each year until 2020. figure 6. total amount of expected deployments 70 vol. 5/2016 deployment of ertms in czech republic no. track distance [km] deployment 1 uzel praha (beroun – praha – benešov u prahy – lysá nad labem) 120 2014-2015 2 české velenice st. hr. rakousko – české budějovice – horní dvořiště st. hr. rakousko 110 2015-2016 3 plzeň – české budějovice 136 2015-2016 4 kolín – havlíčkøuv brod – křižanov – brno 195 2014-2016 5 benešov – votice 20 2015-2016 6 beroun – plzeň – cheb st. hr. německo 169 2014-2016 7 cheb – vojtanov st. hr. německo 20 2015-2016 8 znojmo – šatov st. hr. rakousko 13 2015-2016 9 ústí nad orlicí – lichkov st. hr. polsko 35 2015-2016 10 votice – české budějovice 100 2016-2017 11 hranice na moravě – horní lideč st. hr. slovensko 67 2016-2017 12 ústí nad labem – oldřichov u duchcova/úpořiny – most – karlovy vary – cheb 212 2017-2018 13 pardubice – hradec králové 22 2019-2020 14 zábřeh na moravě – šumperk 13 2019-2020 15 praha – letiště v. havla ruzyně – kladno 31 after 2020 16 brno – přerov 90 after 2020 17 plzeň – domažlice st. hr. německo 72 after 2020 18 velký osek – hradec králové – choceň 96 after 2020 total 1521 table 1. expected track deployment with gms-r [3] no. track distance [km] deployment 1 kolín – břeclav st. hr. rakousko/slovensko 277 2012-2016 2 kralupy nad vltavou – praha – kolín 110 2016-2017 3 st. hr. německo – dolní žleb – kralupy 112 2017-2019 4 petrovice u karviné st. hr. polsko – přerov – břeclav 206 2015-2017 5 praha-uhřiněves – votice 54 2016-2017 6 votice – české budějovice 101 2017-2018 7 česká třebová – přerov 110 2016-2018 8 plzeň – cheb st. hr. německo 117 2016-2018 9 beroun – plzeň 70 2016-2018 10 dětmarovice – mosty u jablunkova st. hr. slovensko 53 2017-2019 11 české velenice st. hr. rakousko – české budějovice – horní dvořiště st. hr. rakousko 112 2017-2019 12 ústí nad orlicí – lichkov st. hr. polsko 37 2018-2020 13 kolín – nymburk – mělník – děčín východ – děčín-prostřední žleb 159 after 2020 14 kolín – havlíčkøuv brod – brno 200 after 2020 15 praha – lysá nad labem 35 after 2020 16 praha – letiště v. havla ruzyně – kladno 31 after 2020 17 praha – beroun 43 after 2020 18 plzeň – domažlice st. hr. německo 72 after 2020 19 pardubice – hradec králové 22 after 2020 20 plzeň – české budějovice 136 aftre 2020 21 brno – přerov 90 after 2020 22 hranice na moravě – horní lideč st. hr. slovensko 67 after 2020 23 uzel praha 40 after 2020 24 český těšín – ostrava-svinov 41 after 2020 25 velký osek – hradeck králové – choceň 96 after 2020 26 cheb – karlovy vary – chomutov 111 after 2020 27 ústí nad labem – chomutov, ústí nad labem – bílina 101 after 2020 total 2603 table 2. expected deployment of equipment upgrade ertm/etcs [3] 71 j. vopava, m. janesova, r. kratochvil acta polytechnica ctu proceedings no. track distance [km] deployment 1 kolín – břeclav st. hr. rakousko/slovensko 250 2015-2019 2 kralupy nad vltavou – praha – kolín 100 2016-2019 3 st. hr. německo – dolní žleb – kralupy 100 2017-2019 4 petrovice u karviné st. hr. polsko – přerov – břeclav 130 2015-2020 5 praha-uhřiněves – votice 40 2016-2020 6 votice – české budějovice 50 2017-2020 7 česká třebová – přerov 50 2016-2020 8 plzeň – cheb st. hr. německo 40 2016-2020 9 beroun – plzeň 40 2016-2020 10 dětmarovice – mosty u jablunkova st. hr. slovensko 20 2017-2020 11 české velenice st. hr. rakousko – české budějovice – horní dvořiště st. hr. rakousko 50 2017-2020 12 ústí nad orlicí – lichkov st. hr. polsko 20 2018-2020 13 kolín – nymburk – mělník – děčín východ – děčín-prostřední žleb 100 after 2020 14 kolín – havlíčkøuv brod – brno 100 after 2020 15 praha – lysá nad labem 30 after 2020 16 praha – letiště v. havla ruzyně – kladno 20 after 2020 17 praha – beroun 30 after 2020 18 plzeň – domažlice st. hr. německo 40 after 2020 19 pardubice – hradec králové 20 after 2020 20 plzeň – české budějovice 50 after 2020 21 brno – přerov 60 after 2020 22 hranice na moravě – horní lideč st. hr. slovensko 30 after 2020 23 uzel praha 20 after 2020 24 český těšín – ostrava-svinov 20 after 2020 25 velký osek – hradec králové – choceň 40 after 2020 26 cheb – karlovy vary – chomutov 50 after 2020 27 ústí nad labem – chomutov, ústí nad labem – bílina 50 after 2020 total 1550 table 3. expected track deployment of etcs l2 [3] after 2020 will be deployed around 250 km tracks and 150 vehicles every year until 2026. strategies transition from the current national system vz ss to european interoperable system class a – etcs consists a combination of investment in track equipment and train equipment. in the first phase will be equipped only international traffic trains. during migration will vehicles equipped with etcs parallel with vz ls that allow operation of both systems together. etcs should be fully migrated for first and second rail corridor and the track přerov – česká třebová at the latest to 2030. after the migration period will be completely decommissioned national systems vz ls and fully deployed etcs system. in some cases can be deployed new special sections equipped solely with etcs, which will be designed to operate only vehicles equipped with etcs even before end of migration period. i.e. a railway connection václav havel airport prague and prague center. 3. deployment of ertms in other european countries unification of europe requires to all transport systems operates without restrictions and time losses, especially when crossing borders. the deployment started in most of european union countries. following countries started participation on deployment of the ertms standard: austria, belgium, bulgaria, croatia, czech republic, denmark, finland, france, germany, greece, hungary, italy, luxembourg, netherlands, poland, romania, slovakia, slovenia, spain, sweden and united kingdom. the figure 7 displays amount of kilometers contracted tracks for each country. the figure 8 displays covering of the ertms divided by level of tracks. the figure 9 displays amount of deployed on-board equipment of vehicles for each country. the figure 10 displays levels of on-board equipment implemented in vehicles. 72 vol. 5/2016 deployment of ertms in czech republic figure 7. ertms contracted tracks [km] in eu figure 8. tracks in eu separated by level 4. conclusions by 2020 the gsm-r will cover major sections of the rail freight corridors in the czech republic. the system will be further deployed progressively on all 3700 km of national routes in order to create comprehensive operational tracks. etcs deployment priority by 2020 is rail freight corridor 7, partially corridors 8, 9, 5 and the track přerov – česká třebová. by 2020, it is necessary to ensure the development of the etcs about 1350 kilometers of tracks and 890 vehicles, ie. for additional six years, around 250 km of tracks and 150 vehicles every year, but this is conditional on the timely implementation of the modernization infrastructure. this requires the necessary financial resources for the trackside and on-board ertms and creating mofigure 9. global ertms vehicles contracted in europe figure 10. vehicles in eu separated by level tivational factors and efficient financial support equipping vehicles with on-board equipment of gsm-r and etcs. references [1] the european rail traffic management system, portál ertms [online]. [2016-02-12], http://www.ertms.net/?page_id=42. [2] národní implementační plán ertms 2014-2020, portál acri [online]. [2016-02-12], http://www.acri.cz/uploads/acri-akademie/15-06% 20etcs/2015_cech,%20varadinov.pdf. [3] národní implementační plán ertms, portál ministerstva dopravy [online]. [2016-02-12], http://www.mdcr.cz/nr/rdonlyras/ 7ff4faf1-1ebe-4116-bf90-808ec7306e04/0/ nipertmsnpweb.pdf. 73 http://www.ertms.net/?page_id=42 http://www.acri.cz/uploads/acri-akademie/15-06%20etcs/2015_cech,%20varadinov.pdf http://www.acri.cz/uploads/acri-akademie/15-06%20etcs/2015_cech,%20varadinov.pdf http://www.mdcr.cz/nr/rdonlyras/7ff4faf1-1ebe-4116-bf90-808ec7306e04/0/nipertmsnpweb.pdf http://www.mdcr.cz/nr/rdonlyras/7ff4faf1-1ebe-4116-bf90-808ec7306e04/0/nipertmsnpweb.pdf http://www.mdcr.cz/nr/rdonlyras/7ff4faf1-1ebe-4116-bf90-808ec7306e04/0/nipertmsnpweb.pdf acta polytechnica ctu proceedings 5:69–73, 2016 1 introduction 2 deployment of ertms in czech republic 3 deployment of ertms in other european countries 4 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0094 acta polytechnica ctu proceedings 12:94–98, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app simulation of electric and hybrid vehicles in a vehicle simulator based on a detailed physical model, for the purpose of hmi evaluation d. rozhdestvenskiya, b, ∗, j. fulemb a czech technical university in prague, faculty of transportation science, horská 3, prague, 120 00, czech republic b their ricardo prague s.r.o., palác karlín, thámova 11-13, 186 00, praha 8, czech republic ∗ corresponding author: rozhddmi@fd.cvut.cz abstract. in this article, we propose a software solution to study hmi of electric and hybrid electric vehicles in vehicle simulators. we will start with the description of a development process of a physical model for hev simulation in ignite software and equation-based language modelica. a short introduction to the language, its possibilities, and explanation, why it is more suitable for the development of such models (vehicle powertrain), are presented in the first part of the article. a fusion mechanism of the physical engine with the model by means of fmu (functional mock-up interface)[1] is also described in this part the second part is dedicated to the description of the model constructed in third party software ignite. this model has a detailed calculation of energy consumption and energy flow based on the selected control strategy. the last part of the article describes a possible experiment methodology. keywords: electric vehicle, hybrid vehicle, powertrain simulation, hmi simulation. 1. introduction electric and hybrids vehicles are slowly gaining popularity not only because it can help to reduce co2 emissions and improve the environmental situation in big cities [2], but because of the rapid development in the area of electric motors batteries and fuel cells technologies. during the latest paris motor show (2016), both vw and mercedes benz claimed to have overcome so far minimal acceptable limit of their pure ev (300 km is a minimal range acceptable by the user when considering to switch from conventional cars to ev or hev [3]) almost twice. this can be achieved because of the new ev and charging architecture developed by companies[4]. additionally, the ambitions of the automotive industry to introduce semi-autonomous and, later, fully autonomous vehicles on public roads before 2020, puts an extra urgency into design and evaluation of hmi technics for such vehicles. user vehicle interaction is shifting to a new, so far, unstudied area with familiar knobs and gauges are being substituted with touchscreens, menus, and voice interactions[5]. the notion of "familiar" car is slowly vanishing, putting a new driver (user) in a position where he has to learn how to drive a vehicle all over again if he switches the brand. this become even more crucial if we take into account that car sharing is slowly substituting car ownership[6]. the processes of development, design, and evaluation of hmi is to adapt to this rapidly changing environment. an efficient way of assessment of vehicle interfaces at different stages of their development can be achieved by testing it in a simulated environment. a model of an interactive ev and hev driving simulator with a possibility of rapid adaptation to vehicle type or architecture, with an accurate simulation of the powertrain components developed in modelica language, is hereby suggested. 2. modelica programming language modelica is an open source object-oriented language for modeling of massive, complex systems. it is suited for multi-domain modeling (including electrical, physical, thermal, mechanical and other domains), additionally, it can be used for hardware-in-the-loop simulations and for embedded control systems [7] models in modelica are mathematically described by differential, algebraic and discrete equations. from a user perspective, models are described by schematic object diagrams). the internal part of a component is defined by another schematic or on a lower level, by an equation-based description of the model in modelica syntax. an example of such a model is shown in figure 1. components can be combined in a complex model by drawing connection lines between "connectors". connectors describe the interaction possibilities (electrical pin, mechanical flange, or an input signal. regarding development time for complex models, modelica language is on average, twice as fast as any other technics in proprietary code or in any other 94 http://dx.doi.org/10.14311/app.2017.12.0094 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 simulation of electric and hybrid vehicles in a vehicle simulator figure 1. modelica model architecture (example of schematic diagram and corresponding programming code). figure 2. development time chart. block modeling languages. this is explained by several factors: • a developer does not have to manually derive relationships between inputs and outputs between subsystems models in modelica are mathematically described by differential, algebraic and discrete equations. no particular variable needs to be solved for manually • modelica code has been developed since 1996 so a lot of subsystem components have been already developed and available for public • implementation time and complexity is the same as for any block modeling language all these points are summarized in figure 2. as mentioned above this puts modelica language ahead of others, especially when dealing with simulation of such a complex and multidisciplinary system as hybrid and electric vehicle powertrain. integration of the model in the virtual world creates a new engineering task which can be solved with an application of the fmi (functional mock-up interface standard). 3. function mockup interface functional mock-up interface (fmi) is a standard to support both model exchange and co-simulation of dynamic models developed in different independent tools, its primary goal is to support the exchange of simulation models between suppliers and oems, providing a possibility to protect product knowledge, which could be recovered from these models. this standard is going to be used to perform a simulation of the ev or hev powertrain inside a physical world simulated in ode or bullet physical engine. figure 3. fmi integration concept[1]. 4. model creation and validation a lot of commercial software available on the market are using modelica as a simulation language, and either already have, or are planning to implement fmu import/export routine. one such software is ignite, developed by ricardo software. ignite is 95 d. rozhdestvenskiy, j. fulem acta polytechnica ctu proceedings a physics-based system simulation package focused at complete vehicle system modeling and simulation. it consists of set of libraries containing vehicle and thermal system components such as engine, transmission, controllers, motors and e.t.c. these components allow faster and easier modeling and parameterization of a powertrain system for any type of vehicle, and a simulation of temperature dependent systems, which will allow further possibilities to simulate hvac and battery management of hev and ev. figure 4. model hybridization. creation of a model of hev in ignite is a straightforward process that consists of several steps [8]: (1.) collect all necessary vehicle information (tires, aerodynamics, mass etc.). vehicle manufacturer usually provides this data in technical documentation. (2.) collect powertrain information (gear efficiencies, shifting maps, engine and electric machine maps and battery characteristics). this data can either be obtained from the automaker directly on request or derived from an experimental data mathematically. (3.) build a conventional vehicle model using ignite libraries: • parametrize the model with collected data. • run test cases to validate the model. • turn the conventional model into a hybrid – add hybrid components (electric motor, battery, controller). validation of the model cacan be performed using standard driver cycles, for instance, ndc or ftp72, results obtained from the simulation on the drive cycle should be compared with an experimental data or figure 5. engine map, torque and soc results. analyzed analytically. results of the simulation contain comprehensive data for every model component and provide the possibility to perform this analysis. an example of the validation for a p2 hybrid with 1.4l tsi motor and 25 kw electrical motor is shown in figure 5. the simulation showed an average fuel consumption of 4.3 l/100km, which is equivalent to 100 g/km co2 emissions on nedc drive cycle for this model. these results closely corelate with the experimental date for similar vehicles. this validation was performed during the development stage of the hybridmodelt at ricardo p.l.c.and described in details in [8]. figure 6. engine map results. 5. integration of fmi into the simulator once the model was created and validated, it has to be modified, to perform a simulation with a physical engine. as we are interested in the interactive simulation for the purpose of hmi evaluation, we need to provide the means of control for a driver to interact with the model. at least 3 inputs have to be 96 vol. 12/2017 simulation of electric and hybrid vehicles in a vehicle simulator introduced to provide a longitudinal control of the simulated vehicle to the user. these inputs are: • acceleration demand (acceleration pedal position <0,1>). • break demand (brake pedal position <0,1>). • current vehicle velocity. velocity input provides a feedback of an actual speed of the vehicle simulated in the ode (open dynamic engine) simulation environment, which is calculated there with respect to elevation grade, tire/road contact friction, collisions, bumps etc. as a direct output, an applied wheel torque should be obtained from the model, this torque is applied to the physical model of the vehicle in the ode or bullet physical engine which is simultaneously handling latitudinal behavior. all other losses (aerodynamic, acceleration, internal and rolling losses) are calculated in ignite model directly. for the purpose of hmi simulation, we have to obtain additional outputs from the model to simulate gauges and control element of hev. these inputs are: • state of battery charge soc. • engine speed. • electric machine engine speed. • battery temperature. • ice state. these outputs can then be further processed to be displayed on a virtual gauge or in any other graphical form. the architecture of this simulation environment is shown in figure 7. figure 7. simulation architecture. 6. possible experiments with detailed consumption model after the vehicle model and simulation environment has been created, an experiment itself should be designed. we can look at four possible types of experiments with evs and hevs: • experiments aimed at determination of general user acceptance of hmi and its efficiency. • experiments aimed at study of acceptance and effect of particular hmi element. • experiments aimed at comparing different hmi. • experiments aimed to determine preferred system types of notifications per different kinds of events. even though any of these types of experiments should be adapted to a specific requirement, one may distinguish ten high-level steps essential for any type of experiment. (janssen at al. 2004): (1.) to determine the goals of evaluation in terms of usability, safety, workload, acceptability, time frame. (2.) to describe a system (systems) which will be tested during the experiment. (3.) to define a high-level scenario (scenarios) in terms of road type, and conditions, visibility, traffic type and actors, tasks, and goals. (4.) define the subject testing cohort in terms of amount, age and experience group. (5.) define subjective and objective evaluation parameters and instruments for data acquisition. (6.) develop experimental instruction for both participants and experimenters. (7.) perform a pilot experiment test and data evaluation. (8.) carry out the experiment. (9.) analyze data. (10.) produce summary report and draw a conclusion. 7. conclusion in this article, we presented a methodology for development and integration of high fidelity powertrain model of ev or hev into three-dimensional virtual simulation environment. this model provides a comprehensive and precise data results to simulate next generation hmi for ev or hev including gauges and functions (such as power-in-out gauges, predictive navigation and others). these models are written in modelica programming language and coupled with ode or bullet physical world with the help of fmi. this approach allows a rapid development and validation of models, and a quick way of switching between them. references [1] t. b. at all. the functional mockup interface for tool independent exchange of simulation models. in proceedings of the 8th international modelica conference. [2] d. rozhdestvenskiy, p. bouchner, a. mashko, et al. dynamic human-machine interface for electrical vehicle design guidelines. smart cities symposium prague 2015, isbn 978-1-4673-6727-1. 97 d. rozhdestvenskiy, j. fulem acta polytechnica ctu proceedings [3] t. smith, h. vardhan, b. malet, d. mingay. are we there yet. thoughts on in-car hmi. 2014. [4] g. e. outlook2016. beyond one million electric cars. international energy agency, electric vehicle initiative. [5] s. heath, p. sant, b. allen. do you feel lucky? why current range estimation methods are holding back. ev adoption hybrid and electric vehicles conference 2013 (hevc 2013), iet year: 2013. [6] p. b. et al. professional vehicle drivers training and performance measurements with use of advanced driving simulator laboratory. 1.ed., prague: ctu in prague, faculty of transportation sciences, department of vehicles, 2013, 1st, electronic edition. isbn 978-80-01-05421-5. [7] j. tobolar, m. otter, t. bunte. modelling of vehicle powertrains with the modelica powertrain library. in: dynamisches gesamtsystemverhalten von fahrzeugantrieben. [8] j. fulem. ignite & wave-rt powertrain model connected to ipg virtual test driving platform. in: ricardo software usercon europe. 98 acta polytechnica ctu proceedings 12:94–98, 2017 1 introduction 2 modelica programming language 3 function mockup interface 4 model creation and validation 5 integration of fmi into the simulator 6 possible experiments with detailed consumption model 7 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0093 acta polytechnica ctu proceedings 13:93–96, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app long-time investigation of cement composite material with micronized waste marble powder: dynamic modules zdeněk prošeka, b, ∗, jaroslav topiča a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b university centre for energy efficient buildings, czech technical university in prague, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: zdenek.prosek@fsv.cvut.cz abstract. this article focus on “blended cement”. the blended cement was created by using waste marble powder (wmp) as a partial replacement for cement. we investigated the influence of wmp on the developing of the dynamic modulus of elasticity and the dynamic shear modulus in time. four different cement composites with wmp as a partial replacement for cement were studied (5, 10, 15 and 50 wt. %) together with reference samples. dynamic modulus of elasticity was monitored during the first 377 days since manufacture by use of non-destructive testing (resonance method). the results showed that wmp in a small amount had a no effect on the dynamic modulus of elasticity and the dynamic shear modulus. keywords: micronized marble powder, cement composites, mechanical properties, long-time investigation. 1. introduction the recycled materials are usually used as an option to non-renewable resources. the problem is mainly variability of these materials, for this cause, is mainly applied to transport infrastructure [1]. examples of the use of waste or recycled materials were spelled out in the study, e.g. [2–4]. this article differs from another primarily utilizing waste from marble industry, directly waste marble sludge. the waste marble sludge generated during the cutting block, or during the finishing process making an environmental problem. the generation of waste is making environmental problems as the valuable national wealth is getting wasted mainly due to lack of management and lack of technology, air contamination, water contamination and loss of vegetation and fauna [5]. the amount of this waste was calculated for the section of the shaq el thoaban, egypt to one million tons per year [6] and for the section of estremoz anticline, alentejo, portugal to 600 thousand tons per year [7]. both areas have approximately the same production of marble stone, but the technical, operational and management practices in the industry are less developed in area shaq el thoaban and they have contributed significantly to greater waste generation. from all the waste produced from marble enterprise, only 10 to 20 % is used in construction. the remainder is refused and accumulated in mounds marble waste close to the extraction quarries [8, 9]. there are some articles that deal with utilized the marble sludge as the replacement for cement in producing the blended cement. the first study found that the addition of the marble sludge to the cement composite is changed crystals portlandite [10]. the second study showed the morphology of the internal structures of cement pastes with and without marble sludge is identical [11]. in all case, marble sludge in a small amount had a good influence on mechanical properties [12, 13]. grains of marble sludge clump together into lumps and thus grains of marble sludge are not homogeneously distributed in the mixture. new option to reduce or restrict this effect is the use of high-speed milling from lavaris ltd. the resulting output is waste marble powder (wmp) and it has better mechanical properties than marble sludge [14]. high-speed milling method separates all the grains of marble sludge. the high-speed milling method was successfully applied for micronized of recycled concrete [15]. 2. materials and samples the cement used was a commercial cem i 42.5r (radotín, czech republic) ordinary portland cement. the wmp used was obtained from the processing of limestone marble in the czech republic as unsorted waste (beroun, jež ltd.), which was micronized using a high-speed mill from lavaris ltd. (libčice, czech republic). the micronized marble powder with fraction 0–40 mm had an average grain and mean grain size of 9.74 and 4.59 microns, respectively. the testing mixtures were designed to keep the identical workability as the fresh reference mixture. the consistency of the mixtures was determined by workability slump test 20 (d20), (table 1). the water ratio is determined as water to cement and marble pow93 http://dx.doi.org/10.14311/app.2017.13.0093 http://ojs.cvut.cz/ojs/index.php/app zdeněk prošek, jaroslav topič acta polytechnica ctu proceedings set/ material cem i wmp water flow expansion bulk density shrinkage [wt. %] [wt. %] ratio [-] test d20 [mm] [kg.m−3] [%] ref. wmp_0 100 0 0.35 190 1919 ± 5 0.19 ± 0.06 wmp_5 95 5 0.32 190 1915 ± 6 0.27 ± 0.11 wmp_10 90 10 0.315 190 1920 ± 3 0.14 ± 0.05 wmp_15 85 15 0.31 190 1879 ± 6 0.21 ± 0.07 wmp_50 50 50 0.305 190 1760 ± 8 0.45 ± 0.16 table 1. composition of the individual set/mixtures. figure 1. experimental set-up for the resonance method. der weight ratio. cement was replaced by the waste micronized marble powder by the amount of 5, 10, 15 and 50 wt. percent (table 1). mixtures made according to table 1 were cast to the rectangular molds with dimensions 40×40×160 mm. all the rectangular molds were measured their dimensions before the mixture was implemented to the mold. these values were applied for determine shrinkage. after casting, these specimens were kept in the molds for 24 h at the laboratory with the temperature 22 ◦c. each set consisted of 6 specimens. after demolding, these specimens were deposited in a laboratory environment at 22 ± 1 ◦c and relative humidity 50 ± 2 %. 3. experimental methods the dynamic modulus of elasticity and the dynamic shear modulus were observed using the resonance method, based on determinations of the natural frequency measured on specimens. resonance method is a non-destructive test. the last measurement was made after 377 days. the resonance method was carried out using the measurement line from brüel & kjær, consisting of a measurement station type 3560-b-120, type 4519003 acceleration transducers, an 8206 impact hammer type and a control computer with software pulse labshop (figure 1.). the resonance method requires information about the dimensions and the weight of each sample earlier than testing was applied. the response signals were transformed from the time to the frequency using the fast fourier transform. the frequency response function (frf) was determined as the ratio of the response and the excitation force in the frequency. the proper basic natural frequencies were evaluated from the frfs. then the dynamic modulus of elasticity and the dynamic shear modulus were evaluated based on the basic longitudinal natural frequency of the samples and the basic torsional natural frequency of the sample, respectively [16, 17]. 4. result and discussion figure 2 shows the evolution of the bulk density, where the free water evaporates over time, thereby changing the bulk density of the samples. figure 3 and 4 shows the evolution of the dynamic shear modulus and the dynamic modulus of elasticity, respectively over time. all of the samples exhibit the steepest increase of the modulus of elasticity in the first day. after the first day, the cement composite rapidly lost the value of dynamic modulus of elasticity. this was due to loss of weight of the sample by natural drying.this phenomenon can be explained by gauging water evaporates from the sample and changes the inner environment in which the wave from the pulse method spreads. the values were stabilized between 7 and 377 days. the curves got smaller decline modulus between 7 and 377 days. there is a noticeable relation between the value of modulus of elasticity and the amount of wmp used in the mixture. the highest modulus had the reference mixture wmp_0 , namely 16.8 gpa for 28 days old samples and 16.0 gpa fot 377 days old samples in the case of the dynamic modulus of elasticity and the dynamic shear modulus was 7.3 gpa for 28 days old samples and 7.1 for 377 days old samples. the lowest value of modulus had the mixture with 50 wt. % of wmp, namely 14.6 gpa for 28 days old samples and 13.9 gpa for 377 days old samples in the case of the dynamic modulus of elasticity and the dynamic shear modulus was 5.9 gpa for 28 and 377 days old samples. 5. conclusions this work is focused on the mechanical properties of the cement composites composed of portland cement and different amounts of waste marble powder. based on the results it can be concluded that: • all of the samples exhibit the steepest increase of the modulus in the first day. 94 vol. 13/2017 long-time investigation of cement composite ... figure 2. time dependence of bulk density (with standard deviation). figure 3. time dependence of dynamic shear modulus (with standard deviation). figure 4. time dependence of dynamic modulus of elasticity (with standard deviation). 95 zdeněk prošek, jaroslav topič acta polytechnica ctu proceedings • after the first day, the cement composite rapidly lost the value of modulus, due to loss of weight by gauging water. • the values of the dynamic modulus of elasticity and the dynamic shear modulus were stabilized between 7 and 377 days. the curves got smaller decline modulus between 7 and 377 days. acknowledgements this outcome was supported by the czech technical university in prague under no. sgs16/201/ohk1/3t/11 and no. sgs17/060/ohk1/1t/11 and by the ministry of education, youth and sports within national sustainability programme i, project no. lo1605. references [1] p. tesárek, k. šeps, j. topič, et al. influence of recycled materials on resulting mechanical properties of cement composites. in modern methods of experimental and computational investigations in area of construction, vol. 825 of applied mechanics and materials, pp. 53–56. trans tech publications, 2016. doi:10.4028/www.scientific.net/amm.825.53. [2] v. corinaldesi, g. moriconi. behaviour of cementitious mortars containing different kinds of recycled aggregate. construction and building materials 23(1):289 – 294, 2009. doi:10.1016/j.conbuildmat.2007.12.006. [3] j. topič, z. prošek, j. hrůza. mechanical properties of the cement paste with fine ground recycled concrete depending on the microstructure. in modern methods of experimental and computational investigations in area of construction ii, vol. 1144 of advanced materials research, pp. 22–27. trans tech publications, 2017. doi:10.4028/www.scientific.net/amr.1144.22. [4] z. prošek, a. osvald, j. topič, et al. influence of recycled concrete composition on its elastic stiffness. in key engineering materials, vol. 677, pp. 288–291. trans tech publ, 2016. [5] m. el-gammal, m. ibrahim, e. badr, et al. health risk assessment of marble dust at marble workshops. nature and science 9(11):144–154, 2011. [6] k. k. alzboon, k. n. mahasneh. effect of using stone cutting waste on the compression strength and slump characteristics of concrete. international journal of environmental science and engineering 1(4):167–172, 2009. [7] l. lopes, r. martins. marbles from portugal. naturstein, online publication, naturalstone-online com 2012. [8] s. shahidan, i. isham, n. jamaluddin. a review on waste minimization by adopting in self compacting concrete. in matec web of conferences, vol. 47. edp sciences, 2016. [9] z. prošek, j. trejbal, j. topič, et al. utilization of the waste from the marble industry for application in transport infrastructure: mechanical properties of cement pastes. in iop conference series: materials science and engineering, vol. 236, p. 012092. iop publishing, 2017. [10] a. a. aliabdo, a. e. m. a. elmoaty, e. m. auda. re-use of waste marble dust in the production of cement and concrete. construction and building materials 50:28–41, 2014. [11] a. ergün. effects of the usage of diatomite and waste marble powder as partial replacement of cement on the mechanical properties of concrete. construction and building materials 25(2):806–812, 2011. [12] b. demirel. the effect of the using waste marble dust as fine sand on the mechanical properties of the concrete. international journal of physical sciences 5(9):1372–1380, 2010. [13] g. menéndez, v. bonavetti, e. irassar. strength development of ternary blended cement with limestone filler and blast-furnace slag. cement and concrete composites 25(1):61–67, 2003. [14] k. šeps, j. topič, z. prošek. the effect of micronized waste marble powder as partial replacement for cement on resulting mechanical properties of cement pastes. in modern methods of experimental and computational investigations in area of construction ii, vol. 1144 of advanced materials research, pp. 54–58. trans tech publications, 2017. doi:10.4028/www.scientific.net/amr.1144.54. [15] z. prošek, l. hlubocký. mechanical properties of cement composite with material based on waste marble powder and crushed limestone. in modern methods of experimental and computational investigations in area of construction ii, vol. 1144 of advanced materials research, pp. 9–13. trans tech publications, 2017. doi:10.4028/www.scientific.net/amr.1144.9. [16] j. topič, z. prošek, k. indrová, et al. effect of pva modification on properties of cement composites. acta polytechnica 55(1):64–75, 2015. [17] z. prošek, k. šeps, j. topič. the effect of micronized waste marble powder as partial replacement for cement on resulting mechanical properties of cement pastes. in advanced materials research, vol. 1144, pp. 54–58. trans tech publ, 2017. 96 http://dx.doi.org/10.4028/www.scientific.net/amm.825.53 http://dx.doi.org/10.1016/j.conbuildmat.2007.12.006 http://dx.doi.org/10.4028/www.scientific.net/amr.1144.22 http://dx.doi.org/10.4028/www.scientific.net/amr.1144.54 http://dx.doi.org/10.4028/www.scientific.net/amr.1144.9 acta polytechnica ctu proceedings 13:94–97, 2017 1 introduction 2 materials and samples 3 experimental methods 4 result and discussion 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0071 acta polytechnica ctu proceedings 13:71–74, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app using homogenization and nanoindentation for microfibril angle determination of spruce lucie kucíkováa, ∗, vladimír hrbeka, b, jan vorela, michal šejnohaa a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b institute of theoretical and applied mechanics as cr, v. v. i., prosecká 809/76, 190 00 prague 9, czech republic ∗ corresponding author: lucie.kucikova@fsv.cvut.cz abstract. this paper is concerned with the evaluation of microfibril angle of spruce. the microfibril angle is defined as an inclination of microfibrils from the longitudinal axis, the direction of lumens. it is well known and further supported by the present study that the microfibril angle has a great influence on the final mechanical properties of wood. this angle could be measured either directly using, e.g. polarisation microscopy, x-ray diffraction, infrared spectroscopy, or indirectly, as used in this study, by employing the nanoindentation measurements. therein, the measured indentation modulus is compared with that obtained numerically using the anisotropic theory of indentation. the latter one depends on the entire stiffness matrix derived through homogenization and the searched microfibril angle. in view of the cell wall microstructure, the effective cell properties were found using the two-step micromechanical homogenization adopting both the self-consistent and mori-tanaka methods. keywords: spruce, nanoindentation, microfibril angle, homogenization. 1. introduction wood is an organic material with highly irregular microstructure on multiple scales. wood is basically divided into softwood and hardwood. in this paper, the norway spruce (picea abies), which belongs to softwood (conifers), was examined. tree stem divides into several components, where the xylem (wood mass) is covering the most part and is the main object of interest. a layered structure appears in the xylem, where earlywood is changing with latewood. one layer of the earlywood and one of the latewood make up the annual ring together. for earlywood, thinner cell walls and larger lumens (ellipsoidal cavity inside the cell) are typical, whereas thicker walls and very thin lumens are typical for latewood. because of this difference, both layers are examined separately. the basic wood constituents are tracheids comprising approximately 94% of the total volume [1]. tracheids are long closed tube-like cells, with almost rectangular cross section, converging on their tips, oriented mainly longitudinally. therefore, the cell wall is the main supporting constituent. the lumens are considered as voids, which are mainly used for transport of water and nutrients. the cell wall could be considered as multilayer laminate, where each lamella corresponds to one layer. furthermore, each layer could be described as a fiber composite. individual layers are basically divided into the warty layer, the secondary wall, the primary wall and middle lamella (sometimes called intercellular layer), described outwards. for more details about ultrastructure of the cell wall see [1–3]. the major part of the cell wall is covered by the middle layer of the secondary wall (s2 layer). the main load-bearing part of the cell wall is the cellulose in the form of microfibrils. microfibrils run helically along the cell wall and the inclination from the longitudinal axis is defined as the microfibril angle (mfa). it is confirmed by many studies that mfa has a great impact on final mechanical properties of the wood. the mfa can be measured directly, see [4]. another possible approach, introduced in this paper, is mfa estimation using comparison between data from the up-scale homogenization and the results of nanoindentation of the cell wall. 2. nanoindentation nanoindentation is nowadays the most commonly applied measurement method to determine material properties on micro and nano scale. this method is based on the contact mechanics with limitations for some types of material responses e.g. viscoelastic solids. the method was developed for measuring hardness and elastic modulus from load-displacement data, obtained during one cycle of loading and unloading by using sharp, self-similar indenters, such as the berkovich triangular pyramid. the basic principle of this method is pushing the small hard tip with known geometry and material properties into a material with unknown properties, with simultaneous recording of the load and indentation depth relative to the initial undeformed surface. the basic output of the nanoindentation measurement is the load-displacement curve, depicted in figure 1, where pmax is the maximum 71 http://dx.doi.org/10.14311/app.2017.13.0071 http://ojs.cvut.cz/ojs/index.php/app l. kucíková, v. hrbek, j. vorel, m. šejnoha acta polytechnica ctu proceedings figure 1. load-displacement curve [5]. figure 2. surface scan (25×25 µm) of earlywood before (left) and after (right) nanoindentation. load, hmax is the maximum displacement, s = dp/dh is the elastic unloading stiffness defined as the slope of the upper part of the unloading curve and hf is the final depth – the permanent penetration depth after the indenter is fully unloaded. in this model, the elastic and plastic deformation during loading is assumed, whereas during unloading only elastic displacements are recovered. for more information about nanoindentation and its principles see [5]. small samples of the norway spruce were embedded in an epoxy resin to stabilize the porous microstructure of the wood. next, they were cut into slices, grounded and polished to obtain as smooth surface as possible. roughness of the sample surface significantly affects the final results of nanoindentation. the measurement was performed on a hysitron tl750 equipped with berkovich diamond indenter. the indents were located to the s2 layer of various cells, with different age, position in a stem both longitudinal and horizontal etc., using click script device. due to the viscoelastic response of wood, the trapezoidal load function with holding part was used. after the measurement, it was necessary to control location of the indents, using surface scans, and exclude wrong positions. the example of the surface scan is shown in figure 2, where the indents on the bottom part are evidently on the wrong position. also curves with wrong shape, caused by surrounder [gp a] coeff. of variation earlywood 12.88 0.18 latewood 13.23 0.19 overall 13.06 0.19 table 1. indentation moduli. ing conditions (temperature, vibrations etc.), surface roughness etc., should be excluded too. after disregarding all wrong load displacement curves, data evaluation can be performed. the standard output of the measurement comprises indentation modulus er, projected area a, maximum load pmax, contact depth hc etc. these values are computed automatically by software implemented in the measurement device. in this case the samples were prepared in a way the indentation direction is practically identical to the longitudinal axis of the cell. from that, the indentation modulus er corresponds to the axial modulus of the cell wall. the normal distribution was used for the data evaluation, where the final number of data is 761 for earlywood, 738 for latewood and 1499 altogether. the resulting means of the indentation moduli are summarized in table 1. clearly, the difference between earlywood and latewood properties is insignificant as was also suggested in [6]. 3. homogenization wood is generally anisotropic material with more or less random microstructure displayed on several scales. this fact promotes micromechanical homogenization as a suitable approach to obtain mechanical properties of the cell wall. for our purpose, only two-scale homogenization was used, following method described in [7]. the homogenization starts at the level called polymer network, with a characteristic length of a representative volume element (rve) in the range of 8–20 nm. two variants were used for the first step. the first one comprises two phases (lignin and hemicellulose), neglecting extractives and considering the dried state of the wood (without water). whereas the second one comprises three phases (lignin, hemicellulose and water+extractives). it is considered that extractives are dissolved in a water and from that it is considered as one phase. in this case the moisture content was set to 0.1, temperature 20°c, relative humidity 0.5 and dry wood density 430 kg/m3. according to authors, all phases are treated on the same footing, from that the self-consistent method seems to be a suitable method. lignin was set as a matrix and the rest of the phases as spherical inclusions. all of the phases are considered as isotropic materials. the input values of all phases and corresponding volume fractions were taken from [7]. the second step provides mechanical properties of the cell wall, which are important for the mfa computation. the cell wall thickness is about 1.5 µm 72 vol. 13/2017 microfibril angle determination of spruce variant 1 variant 2 [gp a] ea et ea et polymer network 6.89 6.89 3.59 3.59 cell wall 44.26 9.34 36.21 5.33 table 2. axial and transverse young moduli of cell wall from two-step micromechanical homogenization. figure 3. angle dependency of the indentation modulus. and 5 µm in case of earlywood and latewood, respectively [1]. the computational procedure is the same for both variants. this step comprises three phases: polymer network, which is set as a matrix, crystalline and amorphous cellulose, both considered as cylindrical inclusions. all inclusions are parallel to the longitudinal axis, which corresponds to zero microfibril angle. following the cell wall fibrous composition, the mori-tanaka method appears to be a suitable choice for this level. both the self-consistent and mori-tanaka micromechanical models are further described in [8]. the results for both variants are summarized in table 2. it is evident that the difference caused by neglecting extractives and water is significant. 4. mfa evaluation this section offers a simplified approach to the derivation of mfa by combining anisotropic theory of nanoindentation and classical micromechanics described in the previous section. by comparing the results of nanoindentation in table 1 and the theoretical predictions from micromechanical homogenization in table 2, although not fully compatible, it becomes clear that the homogenization provides considerably higher values. this discrepancy is caused, among other factors, by the microfibril angle, which affects the final value of longitudinal modulus. in [9] the authors describe the longitudinal modulus ea as a function of mfa and introduce the loading angle, which depends on mfa. as it was mentioned in section 3, the resulting effective matrix is defined for zero microfibril angle. to obtain the effective properties at cell wall level assuming a non-zero mfa, the matrix transformation must be performed. adopting the anisotropic variant 1 variant 2 mfa [°] cov mfa [°] cov earlywood 41.34 0.32 19.01 0.49 latewood 39.69 0.36 19.10 0.48 overall 40.53 0.34 19.05 0.49 table 3. mean values of the mfa. (cov = coefficient of variation). theory of indention presented in [10] we obtain the indentation modulus as a function of the rotated homogenized stiffness matrix to be compared with that measured by nanoindentation. matching these two moduli allows us to extract the searched mfa, see also [11, 12] for further reference. the variation of indentation modulus as a function on mfa is plotted in figure 3. this approach allows us to compare data from nanoindentation measurement and rotated values from homogenization. finding the match of the longitudinal moduli from both methods ehom and eindent, respectively, provides an estimation of the microfibril angle. the resulting mean values of the microfibril angle are summarized in table 3. the results of variant 1 are higher than values of variant 2. the mfa computed with input data of variant 2 is in the range given in literature. for example in [2] author states a range of the mfa of the middle layer about 0° – 30°. 5. conclusions this paper was concerned with several specific topics of the material analysis of wood elastic response. firstly, the direct measurement method on micro scale was introduced. thanks to the nanoindentation, we are able to measure the mechanical properties of the cell wall. nevertheless, the method is still under development and several factors are not known yet. unfortunately, the wood is heterogeneous and very variable material. the material properties are changing with growth conditions in case of the living tree and even when the wood is in the form of the timber, they differ with temperature and humidity. this makes the measurement even more difficult. possibly the main task is the sample preparation. moisture absorption and desorption of the wood causes great problems during preparation. also its softness complicates cutting and the thermal degradation limits the oven drying. even the evaluation of the measured data is still under development, especially in case of viscoelastic materials, to which the wood belongs. despite all these factors, we consider that the nanoindentation gives more or less good approximation of the cell wall mechanical properties. the mean values of the indentation modulus are summarized in table 1. in this paper the evaluation of the measured data for an anisotropic half space was employed [10]. another method introduced in this paper is the micromechanical homogenization. following [7] the 73 l. kucíková, v. hrbek, j. vorel, m. šejnoha acta polytechnica ctu proceedings two-step homogenization procedure was performed to obtain effective mechanical properties of the cell wall. there are two variants of the computation. the first one is neglecting the extractives and considers the dry state, whilst the second one includes extractives and initial conditions are set to: moisture content 0.1, temperature 20°c, dry wood density 430 kg/m3. all input arguments were taken from literature [7]. two computational methods were used – the selfconsistent and the mori-tanaka method. for the first step – polymer network comprising two or three phases (depends on the variant) the self-consistent method was employed. due to the fibrous nature of the cell wall, the mori-tanaka method was chosen as a suitable approach for the second step (cell wall). in this case the layered structure of the cell wall was neglected and the resulting properties written in table 2 correspond to the s2 layer of the cell wall. among other factors, the microfibril angle influences the mechanical properties of the cell wall. except for the various methods of the direct measurement [4], the computational approach could be performed. considering the nanoindentation data as real properties of the cell wall with yet unknown microfibril angle and results from homogenization as the real properties with zero mfa, the angle could be computed as only variable using the comparison of nanoindentation data and rotated values from the homogenization. the mfa-indentation modulus dependency is depicted on figure 3 and the final values of mfa are summarized in table 3. the values of the variant 2 correspond to the data given in literature [2]. there are still many inaccuracies in this method of the mfa determination. very helpful could be the verification of the model using the direct measurement of the microfibril angle on the same samples. due to the visco-elastic response of the wood, the different type of the loading during nanoindentation should be performed. even the sample preparation can be improved, using the microtome cutting rather than grinding and polishing. all these improvements and even more are the task of the future research. acknowledgements the financial support provided by the gačr grant no. 1510354s and by the czech technical university in prague within sgs project with the application registered under the no. sgs17/168/ohk1/3t/11 is gratefully acknowledged. references [1] p. o. kettunen. wood: structure and properties. trans tech publications ltd, uetikon-zuerich, 2006. [2] h. e. desch, j. m. dinwoodie. timber: structure, properties, conversion and use. 7th. macmillan press, hampshire, 1996. [3] m. fujita, h. harada. ultrastructure and formation of wood cell wall. in d. n.-s. hon, n. shiraishi (eds.), wood and cellulosic chemistry, pp. 1–49. marcel dekker, inc., new york, 2001. [4] l. donaldson. microfibril angle: measurement, variation and relationships a review. iawa journal 29(4):345–386, 2008. [5] w. oliver, g. pharr. measurement of hardness and elastic modulus by instrumented indentation: advances in understanding and refinements to methodology. journal of materials research 19(1):3–20, jan 2004. [6] l. melzerová, l. kucíková, t. janda, m. šejnoha. estimation of orthotropic mechanical properties of wood based on non-destructive testing. wood research 61(6):861–870, 2016. [7] k. hofstetter, c. hellmich, j. eberhardsteiner. development and experimental validation of a continuum micromechanics model for the elasticity of wood. european journal of mechanics a/solids 24(6):1030–1053, 2005. [8] m. šejnoha, j. zeman. micromechanics in practice. wit press, southampton, boston, 2013. [9] w. gindl, t. schöberl. the significance of the elastic modulus of wood cell walls obtained from nanoindentation measurements. composites part a: applied science and manufacturing 35(11):1345–1349, 2004. [10] j. vlassak, m. ciavarella, j. barber, x. wang. the indentation modulus of elastically anisotropic materials for indenters of arbitrary shape. journal of the mechanics and physics of solids 51(9):1707–1721, 2003. [11] a. jäger, t. bader, k. hofstetter, j. eberhardsteiner. the relation between indentation modulus, microfibril angle, and elastic properties of wood cell walls. composites part a: applied science and manufacturing 42(6):677–685, 2011. [12] m. šejnoha, t. janda, j. vorel, et al. combining homogenization, indentation and bayesian inference in estimating the microfibril angle of spruce. procedia engineering 190:310–317, 2017. 74 acta polytechnica ctu proceedings 13:72–75, 2017 1 introduction 2 nanoindentation 3 homogenization 4 mfa evaluation 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0048 acta polytechnica ctu proceedings 7:48–52, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app parameters affecting the structural analysis of a tunnel structure exposed to fire omid pourana, ∗, reinhard harteb, carsten peterc a structural engineering, institute for statics and dynamics of structures, bergische universität wuppertal, wuppertal, germany b structural engineering, institute for statics and dynamics of structures, bergische universität wuppertal, wuppertal, germany. kraetzig & partner engineering consultants, bochum, germany c imm maidl & maidl engineering consultants, bochum, germany ∗ corresponding author: pouran@uni-wuppertal.de abstract. behaviour of cut-and-cover tunnels exposed to fire should be analysed by using a realistic structural model that takes account of mechanical and thermal effects on the structure. this has been performed with the aid of finite element (fe) software package called sofistik in parallel, for two types of elements as a scope of research project financed by the german bundesanstalt für straßenwesen bast. since the stiffness of the structure at elevated temperatures is highly affected, a realistic model of structural behaviour of the tunnel could be only achieved by considering the nonlinear analysis of the structure. this has been performed for a 2–cell cut and cover tunnel by taking account of simultaneous reduction of stiffness and strength and the time-dependent increasing indirect effects due to axial constraints and temperature gradients induced by elevated temperatures. the thermal analyses have been performed and the effects were implemented into the structural model by the multi-layered strain model. the stress–strain model proposed by en 1992-1-2 is implemented for the elevated temperature. since there was sufficient amount of polypropylene fibres in the concrete mixtures, modelling of spalling was excluded from the analysis. the critical corresponding stresses and material behaviour are compared and interpreted at different time stages. the main parameters affecting the accuracy and convergence of the results of structural analysis for the used model are identified: defining a realistic fire action, using concrete material model fulfilling the requirements of fire situation in tunnels, defining appropriate time intervals for load implementations. these parameters along with other parameters, which influence the results to a lesser degree, are identified and investigated in this paper. keywords: cut-and-cover, tunnel, fire, thermal effects, structural analysis, accidental situation. 1. introduction in the event of fire, structural behaviour of road tunnels are affected by imposed thermal effects to high extent and this might, consequently, result in failure of the structure. this risk can be minimized by means of appropriate fire design, which provides sufficient passive fire protection for the structure. using an advanced method proposed by en 19921-2 [1] to analyse and design such structures, the following three main areas should be taken into account: • determination of fire action for the specified project, • using realistic material model for the concrete exposed to high temperatures, • defining a thermo-mechanical model which takes account of indirect-thermally induced effects corresponding to the specified fire action and the strength reductions due to temperature rise simultaneously. in germany, the nominal fire exposure curve adopted in for tunnel structures is called ztv–ing fire exposure curve [2]. this can be assumed as fire action applied to the compartmented tunnel structure for accidental fire design situation. further, regarding the material behaviour of reinforced concrete structures, spalling of concrete is an important phenomenon. this implies reduction of cross section thickness and, as a consequence, decrease of concrete cover, which may expose the reinforcement directly to the fire. to avoid spalling, in germany, polypropylene fibres (pp fibres) are usually used to improve the concrete for both new road and railway tunnels. the addition of pp fibres, verifiably, reduces the inner pore pressure in the concrete and by this decreases the risk and impact of concrete spalling [3], as various small and large–scale fire tests have shown. the stress– strain curve proposed by en 1992-1-2 [1] for concrete exposed to elevated temperatures is for the concrete without pp fibres and the heating rate between 2 to 50 k · min−1 which occurs at the fire growth stage for building structures. further, the residual strength at cooling phase should be accordingly determined for each specified project condition. therefore, number of tests have been carried out at bergische universität wuppertal to investigate the effects of polypropylene 48 http://dx.doi.org/10.14311/app.2017.7.0048 http://ojs.cvut.cz/ojs/index.php/app vol. 7/2017 parameters affecting structural analysis of tunnel exposed to fire figure 1. nominal fire exposure curve vs ztv-ing fire exposure curve [4]. fibres on material behaviour of concrete at high temperatures. for verification of ultimate limit states (uls) criteria, structural calculations are performed for fire as an accidental design situation. in this case, the indirect thermal effects result from forces due to restrained thermal elongation or rotation dependent on the temperature distribution in the concrete cross section. the superposition of thermal effects in different time steps with the most unfavourable effects from standard load combinations is performed as so called thermo–mechanical analysis. this analysis is based on an iterative nonlinear calculation to determine the internal forces by considering the simultaneous decrease of stiffness of each cross section exposed to elevated temperatures, and considering the indirect thermally– induced effects. the conditions of equilibrium and compatibility are met by taking into account the maximum permissible temperature–dependent strains for reinforcement and concrete given in en 1992-1-2 [1]. to be able to analyse a 2–cell cut-and-cover tunnel exposed to fire, an advanced analysis according to [1] has been performed by taking into account a model with a simplified boundary condition and nominal ztv-ing fire exposure curve [2]. 2. ztv-ing fire exposure curve for tunnel structures in germany the most common fire exposure curve used for building fire structural design is the iso fire exposure curve according to en 1991-1-2 [1] whereas for road tunnels in germany the ztv-ing curve according to [2] has to be applied. the tunnel shall be designed for the fire situation by considering the high possibility of several vehicles being involved in fire ignition and the fire rate of 100 mw [5]. thus, the increase rate of temperature within the first five minutes from the start of ignition (fig. 1) is very high. further, for the long period of 25 minutes, temperature values remain constant, 1200 ◦c, which is completely different from the nominal fire exposure curve having a continuous trend where the ambient temperature reaches 1006 ◦c, gradually, after 90 minutes. 3. effects of polypropylen fibres on concrete behaviour at elevated temperatures the material properties according to en 1992-1-2 [1] are based on the thermo–mechanical tests for the concrete without pp fibres and they correspond to the defined heating rate between 2 to 50 k · min−1. the latter is to be used for the structural calculations for tunnel structures having pp fibres in concrete’s admixtures only conditionally due to the following reasons: • the german fire exposure curve for tunnel structures includes a cooling phase, which means it comprises a growth, a full and a cooling phase. considering the latter, the changes in the strength of concrete in the cooling phase and the residual deformations are not considered • the addition of pp fibres has influence on the thermal and thermomechanical behaviour of concrete through the experimental investigation at bergische universität wuppertal, influences of specific boundary conditions on the thermomechanical properties were investigated. these were conducted with steady and transient tests according to the recommendation of rilem technical committee [6] on the test specimen with the diameter of 60 mm and the height of 180 mm. for the samples, the concrete strength class of c30/37 with pp fibres and without pp fibres were used. a volume of 2 kg · m−3 of pp fibres has been added to the concrete mixtures to prevent spalling. further, tests were carried out with drying boundary condition and moisture content of 3 % at the beginning of test. to perform the compressive test a static compression and tension machine with a maximum compressive force of 250 kn has been used. heating has been performed via a tube furnace with a maximum temperature of 1200 ◦c. two different deformation measuring devices were installed, one on the specimen and the other one on the test equipment. regarding the material behaviour, the total strain suggested in en 1992-1-2 is according to [7] and has the following components: εtot = εth(t) + εσ(σ,σ,t) + εcr(σ,t,t) + εtr(σ,t,t) (1) where εth, thermal strain including shrinkage, εσ, instantaneous stress related strain, εcr, creep strain or time dependent strain and εtr, transient strain, all resulted from transient and steady tests. it should be noted that en 1992-1-2, however, proposes an implicit model. to be able to model the behaviour of concrete by an explicit strain model, the transient creep component should be defined. transient creep of concrete is the creep that occurs during first time heating at a constant rate under load. 49 o. pouran, r. harte, c. peter acta polytechnica ctu proceedings figure 2. comparison of reduction of compressive strength from the tests results with pp fibres with en 1992-1-2 results for the specimen without pp fibres [9]. further for the drying boundary condition [8]: ε t,0,d tr,cr = ε t,σ,d tr,tot − ε t,0,d tr,th − ε t,0,d tr,sh − ε t,σ,d co,el (2) where εt,0,dtr,cr , transient creep for drying concrete during heating, εt,σ,dtr,tot, total measured transient strain, ε t,0,d tr,th , thermal strain (free thermal strain), ε t,0,d tr,sh , drying shrinkage strain and εt,σ,dco,el , elastic strain. for drying concrete heated without load (i.e. σ = 0) ε t,0,d tr,th + ε t,0,d tr,sh = ε t,0,d tr,tot (3) therefore: ε t,0,d tr,cr = ε t,σ,d tr,tot − ε t,0,d tr,tot − ε t,σ,d co,el (4) two main influences, rising from addition of pp fibres to the concrete mixture, are on the compressive strength of concrete evaluated from the steady state tests unstressed during the heating and on total transient strain, εt,σ,dtr,tot, for specimen under different stress levels, α. the latter corresponds to the ratio of constant value of stress imposed during the heating phase to the compressive strength of concrete specimen at the ambient temperature. results of steady tests as shown in figure 2 from tests at 20, 200, 400, 600 and 750 ◦c compared to the test results from en 19921-2 curves represents a good agreement between the compressive strength reduction of concrete specimen with pp fibres with a compressive strength of 63 mpa. further, there is a considerable influence of loading on the reduction of compressive strength. figure 3 shows the values of thermal strains for concrete specimen with pp fibres. the results show that the addition of pp fibres has an influence on the thermal strain course. thermal strains for pp fibres concrete in the range of 200 ◦c to 300 ◦c shows a small amount of deviation compared to the values of en 1992-1-2. further, the maximum value of thermal strains for pp fibres concretes is lower compared to the en 1992-1-2 curve. in range of 200 to 300 ◦c the drying shrinkage and thermal elongation components figure 3. total strain vs. temperature for specimen with pp fibres [9]. are of main role and due to this there is not large deviation in this range. the lower value of maximum thermal strain for pp fibres concrete is due to a smaller crack divisions at pp fibres concrete and composition of less internal stress. 4. thermo-mechanical analysis in order to find the best approximation of heat transfer through the concrete elements, thermal analyses of five concrete samples with different thicknesses of 80, 90, 100, 110, 120 and 130 cm have been performed for 90 minutes. time intervals of 5 minutes, were chosen to read the absolute temperatures after each step. the absolute temperature values corresponding to each time interval are evaluated at different layers across the thickness of concrete. 1d heat transfer is considered by assuming concrete as a homogeneous and isotropic material. fourier’s law of heat conduction is used for the energy transfer in temperature gradients directions with the upper limit of thermal conductivity of concrete, λc = 1.951 w · k−1 · m−1, for the normal concrete corresponding to 10 ◦c according to en 1992-1-2. further, absolute temperature values as a function of time according to ztv-ing fire exposure curve were applied on the exposed edges. the newton’s law of convection with the heat convection coefficients of correspondingly α = 25 and 50 wm2k are used. the convective coefficient of heat transfer is, however, an experimental value and it is not a material constant. however, the comparison of results by the authors shows that using each coefficient influence the temperature fields in the fire growth phase for the fire exposed boundary up to 10 minutes by 5 %. therefore, the effects are negligible in this range. to find the effect of emissivity ε [10], studied two values of emissivity of 0.7 and 0.8 on the absolute temperatures within a concrete element and concluded that for the depth of 10 cm from the exposed surface there was a temperature difference of 10 ◦c was observed and for the deeper layers there is no influence on the absolute values of temperature. thus, stefan–boltzmann law is assumed with the emissivity grade of 0.8. figure 4 shows how temperature values vary as a function of time on 6 layers across the half–thickness of an 80 cm 50 vol. 7/2017 parameters affecting structural analysis of tunnel exposed to fire figure 4. temperature curves for an 80-cm-thick element at different nodes [11]. thick concrete element till 90 minutes. it is to be noted that temperatures on the boundary, corresponding to node 51, tend to increase with a slower rate compared to the temperatures according to ztv-ing curve up to 30 minutes. having thermal analysis performed for different concrete elements, the thermal effects are implemented to the mechanical model. this is performed by combining the multi-layered strain model [12], taking account of strain induced by thermal effects. the latter takes account of the reduced material strength corresponding to thermal and loading effects at each previous time step as initial governing state for the subsequent step of analysis. based on 10-minute time intervals and nonlinear analysis, the stresses according to each time step plotted and the ultimate limit strength verification of the system for the whole duration of 90 minutes is carried out. the critical points at which the plastic hinges might be developed are identified and possibility of failure by developing further plastic hinges. by reaching 30 minutes, as shown in figure 6, the maximum relative bending moment has reached a value of −2514 knm, approximately 73 % greater than a moment imposed just after the ignition of fire (fig. 5). this is due to the simultaneous increase of tensile stresses on the upper part of beam due to thermal induced bending moments and reduced cross sectional properties. the following parameters can affect the result of nonlinear analysis: • time intervals chosen for thermal analysis (minor influence for intervals of 1 to 10 minutes) • time intervals for load implementations (causes unavoidable inaccuracies for time intervals being larger than 2 minutes) • residual force tolerance at nodes (influences the convergence of fe solution and accuracy of effects of actions) mesh generation (influences both convergence and accuracy of results). 5. conclusions to be able to analyse a 2-cell cut-and-cover tunnel structure with a sufficient amount of pp fibres in the concrete mixtures, three main areas with the corresponding parameters, which can influence the results, figure 5. bending moment diagram for t = 0 [4]. figure 6. bending moment diagram for t = 30 [4]. have been investigated. these main areas were defining a realistic fire action, using an appropriate material model and utilizing a thermo-mechanical model that can represent the nonlinear structural behaviour of the structure. the most important parameters according to each area were highlighted and discussed in this paper. acknowledgements the paper has presented some global results from the research project "verification the fire design of cut-andcover tunnel via nonlinear analysis", financed by bast, and the results of test series carried out at the bergische universität wuppertal. the support by bast is gratefully acknowledged. references [1] en 1992-1-2: design of concrete structures part 1-2: general rules structural fire design 2004. [2] b. f. s. bast. zusätzliche technische vertragsbedingungen und richtlinien für ingenieurbauten (ztv-ing). chap. 3, 5. 2013. [3] k. pistol, f. weise, b. meng. polypropylen-fasern in hochleisungbetonen. betonund stahlbetonbau 107(7):476–483, 2012. doi:10.1002/best.201200024. [4] o. pouran, r. harte, c. peter. nonlinear structural analysis of cut-and-cover tunnels exposed to fire. applications of structural fire engineering asfe’15 2015. doi:10.14311/asfe.2015.005). [5] a. schlüter. passive fire protection for tunnels: guidelines, parameters, reality and suitable measures. tunnel pp. 22–23, 2014. [6] recommendation of rilem tc-htc: mechanical concrete properties at high temperatures-modelling and applications. materials and structures pp. 841–583, 2007. doi:10.1617/s11527-007-9285-2. 51 http://dx.doi.org/10.1002/best.201200024 http://dx.doi.org/10.14311/asfe.2015.005) http://dx.doi.org/10.1617/s11527-007-9285-2 o. pouran, r. harte, c. peter acta polytechnica ctu proceedings [7] y. anderberg. mechanical behaviour at fire of concrete and hyperstatic concrete structures. lund institute of technology, lund, 1976. [8] rilem tc 129-mht: test methods for mechanical properties of concrete at high temperatures recommendation: part 7: transient creep for service and accident conditions. materials and structures 31:290–295, 1998. [9] c. peter. tragverhalten von verkehrstunneln im brandfall mit einer innenschale aus pp-faserbeton. bergische universitaet wuppertal, wuppertal, 2016. [10] c. peter, j. knief, j. schreyer, a. piazolla. rechnericher nachweis des baulichen brandschutzes für tunnel in offener bauweise, schlussbericht zum fe projekt 15.0502/2010erb im auftrag des bmvbs/ der bast 2013. [11] o. pouran, r. harte. prüfgutachten hh 02/14-1 zum forschungsprogramm straßenwesen fe 15.0582/2013/frb erprobung des rechnerischen brandschutznachweises für tunnel in offener bauweise 2014. [12] r. harte, w. b. krätzig, s. y. noh, y. s. petryana. on progressive damage phenomena of structures. computational mechanics 25:404–412, 2000. doi:10.1007/s004660050487. 52 http://dx.doi.org/10.1007/s004660050487 acta polytechnica ctu proceedings 7:48–52, 2017 1 introduction 2 ztv-ing fire exposure curve for tunnel structures in germany 3 effects of polypropylen fibres on concrete behaviour at elevated temperatures 4 thermo-mechanical analysis 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0025 acta polytechnica ctu proceedings 15:25–30, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app isogeometric beam element extended for gridshell analysis edita dvořáková∗, bořek patzák department of mechanics, faculty of civil engineering, czech technical university in prague, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: edita.dvorakova@fsv.cvut.cz abstract. the design of gridshell structures is a complicated process, as the resulting shape of structure depends on initial undeformed grid geometry as well as on the history of boundary conditions used to form the final structure. in practice, the physical model is often used to determine the shape of the structure and the initial grid at the same time. introducing isogeometric analysis into a design of gridshells simplifies the design process as the problem can be easily recalculated when initial grid or boundary conditions change and the resulting shape can be immediately illustrated. the presented paper discusses possibilities in isogeometric gridshell modeling and proposes possible solutions of identified problems. keywords: curved beam element, gridshell structure, isogeometric beam element, nurbs. figure 1. realization of a multipurpose gridshell pavilion in třebešov [1]. 1. introduction gridshell is a light-weight structure gaining its strength from a doubly-curved geometry. such a structure is typically made of wood and it is constructed from initially straight grid of laths. the laths are mutually connected by joints allowing only for the in-plane rotation. the final shape of the structure is reached by subsequent introduction of a load or a prescribed displacement at selected nodes. once the desired shape is reached, the joints are tightened in order to fix the in-plane rotation and stiffen the structure. see fig. 1 for an example of a gridshell structure. it is important to design the structural shape properly to reach high low-bearing capacity. the design process of a gridshell can be divided into three main parts: form-finding, analysis, and construction. formfinding is typically performed by an architect who is seeking for the optimal shape of the gridshell taking into the account the aesthetics and structural behavior of the structure. either physical modeling or computer simulation is used for form-finding, nevertheless these methods (and available tools) do not provide knowledge about structural behavior. when the desired shape is defined the structure is assessed in a finite element method (fem) solver, however the analysis considers only the final shape of the structure. there are no available tools which would take into account the history of the construction steps and would allow to assess all the intermediate construction stages. currently the construction itself relies on trial-error methods supported by a human intuition and experience. our goal is to develop a complex tool for analysis assisted form-finding which would allow to find the optimal shape of gridshell and which would help to propose sequence of construction steps. the idea is to use the isogeometric approach which eases incorporation of the analysis within available design tools and which may bring higher accuracy into the analysis of doubly-curved geometry thanks to the more suitable geometry description. the presented paper discusses the advantages of isogeometric gridshell modeling and also points out some issues connected to its application. 2. isogeometric analysis of beams isogeometric analysis has been introduced by thomas hughes in 2005 [2] with the aim to bridge the gap between the computer-aided design (cad) and the finite element analysis (fea). this goal has been achieved by using the same cad basis functions for geometry description and unknowns approximation in the analysis. thus one model can be shared between the design and the analysis. 25 http://dx.doi.org/10.14311/app.2018.15.0025 http://ojs.cvut.cz/ojs/index.php/app edita dvořáková, bořek patzák acta polytechnica ctu proceedings figure 2. description of b-spline (nurbs) finite element geometry. the geometry is modeled by one patch consisting of two knotspans with quadratic nurbs approximation. 2.1. nurbs-based analysis the most wide-spread technology in cad industry are nurbs (non-uniform rational b-splines). the nurbs curve is obtained by linear combination of cartesian coordinates of the control points p using nurbs functions rpi c(ξ) = n∑ i=1 r p i (ξ)pi. (1) each nurbs function is generated by weighting bspline functions npi with a given weight wi associated with control point pi r p i (ξ) = ni,p(ξ)wi∑n j=1 nj,p(ξ)wj . (2) b-spline functions form a special subset of nurbs functions (corresponding to constant wi) and are derived recursively starting with a piecewise constant functions ni,0(ξ) = { 1 if ξi ≤ ξ < ξi+1 0 otherwise , (3) ni,p(ξ) = ξ − ξi ξi+p − ξi ni,p−1(ξ) + ξi+p+1 − ξ ξi+p+1 − ξi+1 ni+1,p−1(ξ), (4) where p is the degree of the b-spline function, ξi is the coordinate of the ith-knot and ξ ∈ 〈0, 1〉 is a parametric coordinate. the computational domain in isogeometric analysis (iga) is firstly divided into patches, which are further divided into knotspans (often referred to as elements in iga). understanding of knotspans is similar to elements in standard fem, nevertheless higher continuity up to cp−1 between knotspans can be achieved naturally using nurbs, unlike c0 in standard fem. moreover, the nurbs basis functions are generally non-iterpolatory. in the knots (points which are dividing patch into knotspans), the continuity can be locally decreased up to c0 by knot multiplication, which is a standard technique in nurbs technology. this procedure imposes the interpolatory behavior of the functions at the particular point. for better understanding of nurbs see fig. 2, where mapping between real and parametric geometry is illustrated. for more information the reader could refer to [3]. 2.2. beam element formulation the formulation of the three-dimensional beam element [4] is derived in local coordinate system (t,n,b), where t,n,b are tangent, normal and binormal vectors and s is the curvilinear coordinate measured along the centerline of the beam. presented formulation is based on timoshenko beam theory and therefore, in three-dimensional space, there are six independent unknowns: tangential displacement ut, normal displacement un, binormal displacement ub and rotations θt, θn, θb (see fig. 3). a strain-displacement matrix b is defined as ε = br, (5) where ε = {εm,γn,γb,χt,χn,χb}t and r = {ut,un,ub,θt,θn,θb}t . b is derived from the following relations for membrane strain εm, shear strains γn and γb, torsional strain χt, and bending strains χn 26 vol. 15/2018 isogeometric beam element extended for gridshell analysis b(s) r(s) t(s)n(s) ut θt θb ub θn un figure 3. local coordinate system and degrees of freedom of a three-dimensional beam element. and χb εm = u′t −κun, χt = θ′t −κθn, γn = κut + u′n − τub −θb, χn = κθt + θ′n − τθb, (6) γb = τun + u′b + θn, χb = τθn + θ′b. curvature κ is given as κ = ∣∣∣∣ ∣∣∣∣d2r(s)ds2 ∣∣∣∣ ∣∣∣∣ (7) and torsion τ τ = dn(s) ds b(s), (8) where r(s) is the position vector. the material stiffness matrix d is given by constitutive relations n = eaεm, mt = gikχt, qn = ganγn, mn = einχn, (9) qb = gabγb, mb = eibχb, where e is the young’s modulus, g is the shear modulus, ik is the torsional moment, in,ib are the moments of inertia, a is the cross-sectional area, ab,an are the reduced shear areas. finally, the element stiffness matrix k is defined as k = ∫ l 0 bt db ds, (10) where the integral is evaluated using gaussian quadrature, which is not exact for nurbs functions, but it has a sufficient accuracy. 2.3. numerical locking well-known problem of numerical locking of timoshenko beam elements in standard fem occurs also 0 0.5 1 x -0.5 0 0.5 q n shear force q 0 0.5 1 x 0 0.1 0.2 0.3 m b bending moment m figure 4. simple gridshell structure modeled by four patches with cubic approximation. joint is modeled by connection of patches. resulting internal forces overlap exact solution. for the isogeometric formulation. the same approximation order for displacements and for rotations leads to the shear strains of higher order than bending moments, while it should be vice-versa. moreover the constant strains cannot be approximated along the entire patch due to the field inconsistency in the shear terms. the reduced integration, popular locking-treatment in standard fem for its computational efficiency, is not convenient for iga, as the optimal reduced integration rule depends on the choice of approximation order and continuity and such a rule is hard to be determined. on the contrary, discrete shear gap (dsg) method [5, 6] or b̄-method [5] can be used independently of the choice of approximation, the latter being used in the presented implementation of the element. the main idea of the b̄-method is to project membrane strain εm and shear strains γn,γb onto a basis of lower order. for example, the contribution of membrane strain εm to the b̄ matrix is b̄m = [ b̄ 1 m b̄ 2 m 0 0 0 0 ] , (11) where for example b̄ 1 mij = ∫ l 0 ñin ′ j ds (12) and where nj is the jth basis function of the original basis and ñi is the ith basis function of the lower order basis. the membrane contribution to stiffness matrix km is km = ea b̄ t mm̃ −1 b̄m (13) where m̃ is the element “mass matrix” calculated in 27 edita dvořáková, bořek patzák acta polytechnica ctu proceedings 0 0.5 1 x -0.5 0 0.5 q n shear force q 0 0.5 1 x -0.2 0 0.2 0.4 m b bending moment m 0 0.5 1 x -1 0 1 q n shear force q 0 0.5 1 x -0.2 0 0.2 0.4 m b bending moment m a) c3-continuity b) c2-continuity 0 0.5 1 x -1 0 1 q n shear force q 0 0.5 1 x -0.2 0 0.2 0.4 m b bending moment m 0 0.5 1 x -0.5 0 0.5 q n shear force q 0 0.5 1 x 0 0.1 0.2 0.3 m b bending moment m c) c1-continuity d) c0-continuity figure 5. resulting internal forces (blue continuous line) and exact solution (red dash-dot line) for different continuity at the position of concentrated force load. the lower order basis m̃ij = ∫ l 0 ñiñj ds. (14) the analogical procedure can be followed also for the shear strains γn and γb. 3. application to gridshells isogeometric formulation of beam element allows to perform the analysis directly within the cad system, however the modeling of gridshells with presented isogeometric beam element formulation is not straightforward. there are two possible approaches to modeling gridshells: a) multiple patches per one lath, b) single continuous patch per lath. 3.1. lath modeled by multiple patches this approach is similar to the use of standard fem, where the joint is usually modeled by inserting a node. in iga, it means to assign at least one patch to each lath segment between two joints leading to the interpolatory degrees of freedom at particular points (joints). this approach is simple and does not require any additional implementation, nevertheless it leads to loss of some advantages of isogeometric formulation, namely loss of higher inter-element continuity along the entire lath and loss of possibility to handle one lath as a single entity, which might be highly convenient for the development of a complex design tool. a simple test has been performed to verify the "multiple-patches" approach. two laths mutually connected in the middle, each modeled by two patches with one knotspan and cubic b-spline approximation, have been subjected to the transverse force placed at the joint. the model and resulting shear force γn and bending moment χb are shown in fig. 4. the results 0 0.5 1 -0.1 0 0.1 n * 0 0.5 1 0 0.05 0.1 n v * figure 6. additional basis functions to support case of concentrated force load on 2d straight beam with nurbs approximation. overlap the exact solution and prove the possibility to use this modeling technique for gridshell analysis. 3.2. lath modeled by a single patch isogeometric approach offers a possibility to model each grid lath by a single patch. such a model requires to add additional constraints to enforce the compatibility of displacements and rotations in joints as the compatibility cannot be achieved by sharing particular degrees of freedom due to the non-interpolatory nature of isogeometric basis functions. a special boundary condition based on the lagrange multipliers method has been implemented to support this approach. the potential energy is enhanced by an additional term yielding the required continuity wint + ∑ k λk(r̄ki − r̄ k j ) = wext, (15) where λk is vector of lagrange multipliers for kth joint, r̄ki and r̄ k j are the vectors of constrained unknowns at the kth joints on the ith and jth beams, respectively. note, that due to the non-interpolatory nature of nurbs basis functions, the vectors r̄ki and r̄ k j have to be expressed as a linear combination of the control points values using nurbs basis functions evaluated 28 vol. 15/2018 isogeometric beam element extended for gridshell analysis 0 0.5 1 x -1 -0.5 0 0.5 1 q n shear force q 0 0.5 1 x 0 0.1 0.2 0.3 m b bending moment m 0 0.5 1 x -2 -1 0 1 2 q n shear force q 0 0.5 1 x 0 0.1 0.2 0.3 0.4 m b bending moment m figure 7. resulting internal forces for simply supported beam subjected to a) a concentrated force, b) a concentrated force and continuous force loading. nurbs approximation enriched by the functions tailored to support case of concentrated force has been used. the results overlap the exact solution. at the joint position. the local coordinate ξi (ξj) of the joint on the beam i (j) has to be calculated according to the parametrization. the physical meaning of lagrange multipliers λ corresponds to the forces and moments required to enforce the compatibility and thus such an approach results in concentrated force and moment loadings at the joints positions. the problem of such a formulation is the inability of nurbs basis functions to represent exact solution with discontinuities in strains and internal forces corresponding to concentrated loadings. this inability is a consequence of the higher continuity of nurbs basis along the entire patch. this problem is documented on the similar test as the one used in section 3.1 (see fig. 4). unlike in previous section, each lath is now modeled by a single patch using just a single span with cubic bspline approximation, which results in c3-continuity at the joint position causing the oscillations in internal forces, see fig. 5a for the results. the continuity can be lowered to c2 by placing a knot at the particular position, the continuity can be further reduced up to c0 by increasing the knot multiplicity. the resulting internal forces for all choices of continuity are shown in fig. 5. the oscillations in the numerical solutions are sufficiently removed only in the case of c0-continuity. note, that the procedure of decreasing the continuity to c0 results in the same model as in sec. 3.1 and thus the advantages of isogeometric analysis are eliminated. the main idea, how to preserve the higher continuity while avoiding oscillations in numerical solutions, is to enrich the nurbs basis by the special basis functions tailored to match the exact solution. in this paper, the derivation of the additional basis functions is illustrated on the case of two-dimensional straight beam subjected to concentrated force. from the knowledge of the exact solutions of bending moment mex and shear force qex under the unit force mex = eiβ′(x), (16) qex = kga (v′(x) + β(x)) , (17) where β is a rotation and v is a transverse displacement, the additional basis functions n∗β, n ∗ v can be derived n∗β = β = ∫ 1 ei mex dx, (18) n∗v = v = ∫ 1 kga qex −β dx. (19) see fig. 6 for the resulting basis functions. the final approximation for the enhanced solution is v(x) = ∑ nivi + n∗vv ∗, (20) β(x) = ∑ niβi + n∗ββ ∗. (21) the integration constants were determined with respect to the boundary conditions of a simply supported beam. the performance of the enhanced approximation has been tested by means of two simple test. firstly, simply supported beam subjected to the concentrated force load was considered, in the second test the concentrated force has been combined with continuous loading. the resulting internal forces for both load cases overlap exact solutions, see fig. 7. 29 edita dvořáková, bořek patzák acta polytechnica ctu proceedings 4. conclusions two possible approaches to gridshell modeling have been presented. the simple fem-like approach, where the joints are modeled by connection of two patches, provides good results, however the advantages of isogeometric approach are lost. in order to keep the higher continuity along the entire lath, it has to be modeled by a single patch. in such a case, the enforcement of continuity of displacements and rotations in joints by lagrange multipliers inevitably leads to application of concentrated forces and moments. it has been shown, that enhancement of nurbs basis functions by a set of the special functions is sufficient to support concentrated loadings. the paper documents the first necessary steps of the development of a complex design tool for gridshells. the presented idea of enhancement of the nurbs basis has to be combined with locking treatment to provide robust element formulation, moreover due to the large deformations, the formulation has to be extended for geometrically non-linear analysis. acknowledgements the financial support of this research by the grant agency of the czech technical university in prague (sgs project no. sgs18/037/ohk1/1t/11) is gratefully acknowledged. references [1] a. vaňek. multipurpose gridshell pavilion in třebešov, czech republic, physical model as a groundwork for realization, harvest super, 2016. [2] t. j. r. hughes, j. a. cottrell, y. bazilevs. isogeometric analysis: cad, finite elements, nurbs, exact geometry and mesh refinement. comput methods appl mech engrg 194:4135–4195, 2005. [3] l. piegl, w. tiller. the nurbs book. springer-verlag berlin heidelberg, new york, 1997. [4] g. zhang, r. alberdi, k. khandelwal. analysis of three-dimensional curved beams using isogeometric approach. engineering structures 117:560–574, 2016. [5] r. bouclier, t. elguedj. locking free isogeometric formulations of curved thick beams. comput methods appl mech engrg 245-246:144–162, 2012. [6] r. echter, m. bischoff. numerical efficiency, locking and unlocking of nurbs finite elements. comput methods appl mech engrg 199:374–382, 2010. 30 acta polytechnica ctu proceedings 15:25–30, 2018 1 introduction 2 isogeometric analysis of beams 2.1 nurbs-based analysis 2.2 beam element formulation 2.3 numerical locking 3 application to gridshells 3.1 lath modeled by multiple patches 3.2 lath modeled by a single patch 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.12.0064 acta polytechnica ctu proceedings 12:64–67, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app subjective methods for assessment of driver drowsiness alina mashko czech technical university in prague, faculty of transportation sciences, konviktská 20, prague 1, czech republic correspondence: mashkali@fd.cvut.cz abstract. the paper deals with the issue of fatigue and sleepiness behind the wheel, which for a long time has been of vital importance for the research in the area of driver-car interaction safety. numerous experiments on car simulators with diverse measurements to observe human behavior have been performed at the laboratories of the faculty of the authors. the paper provides analysis and an overview and assessment of the subjective (self-rating and observer rating) methods for observation of driver behavior and the detection of critical behavior in sleep deprived drivers using the developed subjective rating scales. keywords: driver drowsiness, subjective methods, observer rated sleepiness. 1. introduction driving sleepy, drowsy or fatigued at a high extent affects road traffic safety. the research by american centers for disease control and prevention (cdc) reports that 7500 fatal accidents happened in years 2011-2012 due to the drowsy factor [1]. per aaa foundation for traffic safety, 21% of drowsy drivers were involved in fatal traffic accidents in the usa, which is 4,5% higher as compared to the results of the preceding years [2]. similar involvement rates (20% of accidents due to sleepiness at the wheel) have been reported by the australian transport accident commission [3]. two-thirds of the population (71%) are said to be sleeping less than 8 hours per day. the most vulnerable to sleep deprivation are so called risk group that includes shift workers, young drivers and people with children in household. driving drowsy is dangerous, mainly due to the possible consequent impairments caused by this state. according to the national sleep foundation’s poll in 2000, drowsy drivers tend to feel stressed (42%), be impatient (32%) or drive faster (12%) [4] or not use the seat belt during night driving [5]. known reasons of drowsy driving and inattention are driving at the time of day when circadian rhythms are low, or driving for a long time or sleep loss [6]. 2. definition of fatigued and drowsy state due to the individual behaviour of every person in a particular state as well as because of a variety of reasons for drowsy driving, finding common attributes or indicators in this state is especially difficult. drowsiness and fatigue, per self-reported perceptions, influence performance, physiological state, feeling. it has been found that in a state of fatigue individuals have experienced such feelings as being tired, sleepy or exhausted, not feeling inclined to continue performing the task because of a perceived reduction of efficiency, feeling tired, exhausted, lack of energy as related to impaired of both physical and cognitive functioning, need of restoration of resources [7]. individual perception of one’s drowsiness can be determined with the help of questioning objects, conducting self-evaluation tests and is related to subjective observation. according to [8], the physiological state of a fatigued individual this corresponds to the organism’s muscular, visceral and central nervous system a decrease of capacity or energy to maintain initial activity due to insufficient rest is observed. detection of physiological signs of drowsiness and fatigue is carried out via measurements of the heartbeat rate, brain activity, eye blinking and closure. eye behavior can be observed with the help of electrooculogram (eog) [9, 10], with eye tracker, such as in researches presented in [11–13], based on eye image video analysis [14], etc. a lot of research has been done for exploration of brain activity of sleepy and drowsy drivers, some of the latest are [15–19]. the measurement itself is performed with the help of special electrodes of an eeg device that are placed on the human scalp to read brain wave activity. evaluation of the quality of the task performance is another possible measure of drowsiness and fatigue. fatigue caused by sleep deprivation affects driving performance which can reflect in delayed reaction, inability to maintain constant speed, loose of lateral control [20, 21] etc. measurements can be conducted in real traffic as well as in simulated environment, which is safer. usually, the performance indicators are measured for each object under two conditions: in a state of sleep deprivation and in fresh (control or reference) state. 64 http://dx.doi.org/10.14311/app.2017.12.0064 http://ojs.cvut.cz/ojs/index.php/app vol. 12/2017 subjective methods for assessment of driver drowsiness 3. subjective drowsiness and fatigue subjective indicators of sleepiness level include sleepiness scales, such as epworth sleepiness scale (ess), karolinska sleepiness scale (kss) and stanford sleepiness scale. during these tests short questionnaires are given to subjects who are asked to rate their state according to the suggested scale. the self-rating is usually conducted repeatedly during the testing time with either a time interval, or situations (conditions). this kind of rating allows evaluation of the participants’ awareness of their alertness level. when compared to other driver drowsiness measures, it has been demonstrated that kss defined level of sleepiness in most cases has a curve-linear relation to lateral position and blink duration [22]. another method used in experimental investigation of driver fatigue is observer rated sleepiness (ors). the method quantifies observation of behavior by trained individuals or experts who are rating driver performance based on video recordings of the experiment or in real time monitoring. the method can be used for validation of other measures or as a separate measure of driver fatigue [23]. 4. subjective evaluation during experiments on driver simulator all observations described in this chapter are related to experiments conducted on a full body (skoda superb) fixed driving simulator based on a wide projection angle (210° of driver’s view is provided with 5 screens plus 2 screens for side mirrors. see figure 1 for reference). detailed description of simulator and lab is provided in [24]. figure 1. advanced driver simulator for driver drowsiness research. previously, we have conducted observations of sleep deprived drivers during a 2 hours’ driver simulator measurements [25]. different observers tracked changes in the drivers behavior based on video recordings. the target areas were head, face (facial expression and mimicking, eye behavior) and upper torso (hands, shoulders and body). observers were instructed to document all noticed behavioral patterns out of (but not limited to) the suggested list: frequency of eye blinking, change in duration of eye closures, yawning, face muscles relaxation, lip biting, head tilt, shoulders shaking, hands on the wheel position, body relaxation, change of sitting position. also from the shots of the driving scenery driving behavior (steering, correction of position in a lane, lane departure and accident) were observed. it was possible to notice that blinking frequency was observed after 40th minute of driving. blinking was then gradually substituted by longer eye closures often accompanied by worsening of steering behavior, such as weaving and loose of vehicle control up to accidents. in most drivers head rotations (neck muscles stretching) and head tilt were noticed. this, however, didn’t reflect in the alteration of driving behavior (i.e., steering, speed fluctuation etc.) and was considered as a voluntary type of behavior caused by both sleepiness and fatigue from a long drive taken by subjects with a purpose of finding more comfortable position. all observed subjects have displayed dangerous driver behavior, that in real traffic might lead to accident, mainly starting after 30 minutes of experiment. inability to maintain straight trajectory, weaving and lane departure could be observed. at the laboratory for driver sleepiness research a self-rating sleepiness scale has been developed and is being used to assess driver sleepiness rate by the subject him/herself. the scale is presented in table 1, each of its levels corresponds to change in sleepiness level as perceived by subject. state description as perceived by subject 1 i feel fine/fresh & driving does not make me any problems. 2 i feel drowsy & driving does not make me any problems. 3 i feel drowsy & i notice some problems. 4 i feel very drowsy & i need to excessively concentrate to drive correctly. 5 i experienced ’blackouts’ & losing control over the car. table 1. self-rating sleepiness scale developed at faculty of transportation sciences, ctu in prague. evaluation per this scale will be presented in further research outcomes. the ors method was applied for analysis of video recordings of sleep deprived drivers on simulator. 9 observers (7 male and 2 female) who are students of faculty of transportation sciences, czech technical university conducted evaluation. the observations combined body movements, facial expressions and driving and vehicle behavior. initially developed scale contained three basic levels: (1.) subject is sitting straight behind the wheel with eyes open and both hands on the wheel in 10/2 position, he/she is following the road situation and 65 alina mashko acta polytechnica ctu proceedings figure 2. results for ors evaluation by 9 observers of one sleep deprived subject. figure 3. difference between maximum and minimum rates provided by observers. cluster indicators, vehicle is within the lane and is following straight path, (2.) subject experiences some features of drowsiness, like more frequent blinking and episodes of prolonged blinking, vehicle is experiencing some weaving, correction movements are present in steering, subject sitting position is not straight (head inclinations), subject is performing some self-awakening actions, hands positioned in lower part of the wheel, (3.) episodes of lane departure due to the result of longer eye closures (sleep episodes) steering correction movements are more abrupt, body position is not straight (subject adjusts to be more comfortable for rest position), head inclined to the side, one hand may be off the wheel, or on the bottom of the steering wheel. during the evaluation process, it had been decided to extend it to a 5-scale grade by introducing the half-scores for so called transition states (i.e., giving 1.5 and 2.5 scorers when subjects demonstrated only some features of the state). outcome of such observation is presented in figure 2. there one may track the mean, max and min rates of subject state sampled per 5 minutes of two-hour experiment. every human behavior is individual and people display different behavior under same conditions. besides, the subjective perception of each observer influences the ors data (see the difference chart between rates – figure 3). the same type of behavior may vary per subject in severity and impact that’s why it is not possible to assign same weights to the displays of human behavior. for those reasons, precise definition of a state to correspond the rate score is sometimes compromised. this is confirmed by a study, where results have demonstrated poor correlation of ors to srs (self-reported sleepiness) [26]. it is still possible to track a tendency of an increasing sleepiness rate per suggested scale after the 40th minute of driving, which corresponds to conclusions mentioned in [25]. 5. conclusion the research of driver fatigue covers the applications of many branches of science such as sociology, physiology, neuroscience, engineering. individual human behavior is a challenge in finding a single method for detection of drowsiness. besides, validation of detection methods in real traffic conditions is more complicated because these experiments may cause serious accidents. subjective methods have proved to be reliable for rating of fatigue in laboratory experiments and can be considered as a viable method for further exploration of the problem. self-rating and expert observations have proven to provide an adequate picture of driver fatigue and the tendency of increasing sleepiness during prolonged driving and can be correlated to objective measurements. rigid definition of the states to correspond the scale rate is however crucial for evaluation results. difference in scores given 66 vol. 12/2017 subjective methods for assessment of driver drowsiness by observers to the same behavior is crucial of data validity and could make the results questionable. the current research is aimed at tailoring the suggested evaluation method for its further implementation for analyses of driver sleepiness experiments. acknowledgements this research was supported by grant sgs16/254/ohk2/3t/16 "experimental research of driver fatigue by means of observation visual behavior". references [1] a. wheaton, r. shults, d. chapman, et al. drowsy driving and risk behaviors – 10 states and puerto rico, 2011-2012. morbidity and mortality weekly report, vol. 63, no. 26. july 4, 2014. [2] aaa foundation for traffic safety. acute sleep deprivation and risk of motor vehicle crash involvement, 2016. [3] transport accident commission. fatigue statistics. http://www.tac.vic.gov.au/road-safety/ statistics/summaries/fatigue-statistics. [4] national sleep foundation. 2000 facts and stats. http://drowsydriving.org/about/facts-and-stats. [5] u. d. of transpotation. nhtsa.guidelines to observe and estimate statewide seat belt use at night. 2010. dot hs 811 288. [6] u. d. of transpotation. nhtsa.guidelines to observe and estimate statewide seat belt use at night. 1999. dot hs 808. [7] r. phillips. a review of definitions of fatigue – and a step towards a whole definition. transportation research part f 29 (2015) 48-56. [8] r. soames-job, j. dalziel. defining fatigue as a condition of the organism and distinguishing it from habituation, adaptation and boredom. in p. a. hancock & p. a. desmond (eds.), stress, workload and fatigue. ny: crc press. 2000. [9] m. villamizar, j. suarez, j. villanueva, et al. design and implementation of sleep monitoring system using electrooculographs signals. 9th pan american health care exchanges conference, pahce 2014. [10] n. r.-i. nez, p. meca-calderón, m. garcía-gonzález, et al. drowsiness detection by electrooculogram signal analysis in driving simulator conditions for gold standard signal generation. international conference on biomedical electronics and devices, biodevices 2013. barcelona, spain. [11] r. coetzer, g. hancke. eye detection for a real-time vehicle driver fatigue monitoring system. 2011 ieee intelligent vehicles symposium, iv’11; baden-baden, germany. [12] k. anjali, k. athiramol, v. athira, et al. time nonintrusive monitoring and detection of eye blinking in view of accident prevention due to drowsiness. international conference on circuit, power and computing technologies, 2016. [13] a. rahman, m. sirshar, a. khan. real time drowsiness detection using eye blink monitoring. national software engineering conference (nsec), 2015. [14] j. ahmed, l. jian-ping, k. saeed, a. riaz. eye behavior based drowsiness detection system. 12th international computer conference on wavelet active media technology and information processing (iccwamtip). 2015. [15] g. usha, b. radhika, t. shanmuganathan. eeg-based brain wave sensor to detect drowsiness with eye open. far east journal of electronics and communications. volume 16, march, 2016. [16] r. foong, k. ang, c. quek, et al. an analysis on driver drowsiness based on reaction time and eeg band power. 37th annual international conference of the ieee engineering in medicine and biology society, embc, 2015. [17] t. brown, r. johnson, g. milavetz. identifying periods of drowsy driving using eeg. 57th annual scientific conference of the association for the advancement of automotive medicine, canada, 2013. [18] m. jiřina, p. bouchner, s. novotný. identification of driver’s drowsiness using driving information and eeg. neural network world, 2010. [19] j. faber, j. pěkný, r. pieknik, et al. simultaneous recording of electric and metabolic brain activity. neural network world, 2010. [20] r. v. loon, r. brouwer, m. martens. drowsy drivers’ under-performance in lateral control: how much is too much? using an integrated measure of lateral control to quantify safe lateral driving. accident analysis and prevention. volume 84, 2015. [21] s. james, b. vila. police drowsy driving: predicting fatigue-related performance decay. policing, volume 38, issue 3, 2015. [22] m. ingre, t. akerstedt, b. peters, et al. subjective sleepiness, simulated driving performance and blink duration. journal of sleep research. volume 15, issue 1, 2006. [23] a. anund, c. fors, d. hallvig, et al. observer rated sleepiness and real road driving: an explorative study. plos one, 2013. [24] s. novotný, p. bouchner, j. pěkný, et al. hybrid mirrors for driving simulators – design, construction and experiments. in: computational intelligence, man-machine systems and cybernetics. athens: wseas press, 2007, pp. 310-315. issn 1790-5117. isbn 978-960-6766-21-3. [25] a. mashko, p. bouchner. visual investigation of driver behavior for fatigue detection – driving simulator experiments. recent advances on systems, signals, control, communications and computers. 2015. [26] c. ahlstrom, c. fors, a. anund. eur. transp. res. rev. (2015) 7:38. doi:10.1007/s12544-015-0188-y. 67 http://www.tac.vic.gov.au/road-safety/statistics/summaries/fatigue-statistics http://www.tac.vic.gov.au/road-safety/statistics/summaries/fatigue-statistics http://drowsydriving.org/about/facts-and-stats acta polytechnica ctu proceedings 12:64–67, 2017 1 introduction 2 definition of fatigued and drowsy state 3 subjective drowsiness and fatigue 4 subjective evaluation during experiments on driver simulator 5 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0057 acta polytechnica ctu proceedings 22:57–61, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app monitoring and protection of timber-concrete bridges petr kuklík∗, lukáš velebil, anna gregorová, petr svora czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: kuklik@fsv.cvut.cz abstract. the majority of timber bridges for ordinary road traffic are often made as timber-concrete composite structures. concrete slab is something like as umbrella for timber. the reason for problems in timber-concrete bridges is nearly always poor detailing for durability and neglected maintenance. mainly all aspects of timber protection and maintenance should therefore be considered even during the planning phase of bridge. paper deals with evaluation of timber properties, system of opto-fiber sensors for monitoring and diagnostics of mechanical stress of timber-concrete bridges, and degradation of timber due to uv radiation, atmospheric conditions and biological agents after application of photocatalytic materials on surface. keywords: bridges, concrete, timber, timber-concrete composite, monitoring, durability. 1. introduction timber-concrete composite bridges were first constructed in the united states in the early 20th century. this development was motivated by a shortage of steel that urged builders to use other available structural materials or combinations of such materials [1]. the first reported timber-concrete bridge in the united states dates back to circa 1925. in the 1950s, timberconcrete composite structures started to be applied to bridge construction in other locations, such as australia and new zealand. since the beginning of the 1990s, timber-concrete composite bridges have been increasingly used mainly in some northern and western european countries, such as finland, switzerland, france, germany, and austria. from the beginning of applications different systems and configurations have been used for bridges. in general, these can be grouped in two categories, see fig. 1. in the design and performance of timber-concrete composite bridges, a key issue is the connection system, which has a direct influence on both the serviceability limit states and the ultimate limit states. bridge structures are normally designed for a long service life, 100 years. the reason for decay in timber bridges is nearly always poor detailing for durability, low quality of materials applied and neglected maintenance (a regular check of the whole bridge should be made at intervals not exceeding three to five years). structural timber as a natural product is part of a life cycle of growing and decomposition. due to that is also the best to use of either naturally durable or chemically treated timber. important is also to keep timber constantly dry as is possible. this paper deals with evaluation of properties of structural timber by semi-destructive method similar figure 1. timber-concrete bridge configuration (a) timber beam, (b) timber deck. to evaluation of concrete, new possibilities of monitoring of bridge structure and new chemical preservation on timber surface. 2. evaluation of density and mechanical properties of structural timber one of the semi-destructive methods of timber evaluation in timber structures is based on measurement of material resistance against penetration of an indenter. the damage of a timber structural member is small after the tests. one of the well-known devices is pilodyn 6j forest that penetrates the surface layers of timber element by a short pin with the help of an internal spring. it is possible to determine wood density and even the related mechanical properties with the input of the depth of the pin penetration [2]. wood density is of key importance value of timber. very important also for the use of non-destructive and semi-destructive methods. wood density is determined by several factors, e.g. cell diameter and cell wall thickness, proportions of springwood and summerwood, cellulose and lignin content etc. den57 http://dx.doi.org/10.14311/app.2019.22.0057 http://ojs.cvut.cz/ojs/index.php/app p. kuklík, l. velebil, a. gregorová, p. svora acta polytechnica ctu proceedings sity has a significant influence on wood mechanical properties and the effect of a non-destructive parameter is frequently associated with the density [3]. the depth to which the pin penetrates is indicated on the instrument and is inversely proportional to the wood density. the equations below can be used for estimation of wood density ρ12. the calculations were developed to be used for in situ measurement of timber quality in the existing structures [4] and [5]. they are based on moisture content (mc) measurement, which is an integral part of each testing of wood, and the depth of penetration tp. ρ12 = −0.027102 · tp,12 + 0.727987 (1) tp,12 = tp · (1 − 0.007 · ∆w) (2) ∆w = w − 12 (3) where • ρ12 is density of wood with mc of 12% [g·cm−3], • tp,12 the depth of penetration into wood with mc of 12% [mm], • tp the depth of penetration into wood with measured mc [mm], • w measured mc [%]. not only the strength fm, but also the modulus of elasticity e can be estimated with help of penetration depth tp or more precisely density ρ12. the relationships between the quantities with specific constants empirically determined resulted from the research [4]. fm = −36.13 + 0.156743 · ρ12 (4) estat,05 = −14 441.49 + 55.7834 · ρ12 (5) 3. mechanical stress monitoring of timber-concrete bridges by fbg sensors due to the increasing number of bridge construction faults, there is a need for monitoring their condition, which could reveal the emerging issues in time and warn of imminent violations. one of the innovative methods particularly suitable for long-term continuous monitoring of structural stress is the use of a fiber optic sensor. the technology of fiber bragg grating (fbg) sensor is an optical system used for the same purposes as strain gauge sensors. however, unlike strain gauges, deformation measurements and information transmission are realized by means of light. energy requirements of the system are therefore reduced and the system is significantly more resistant to most potential sources of interference, including electromagnetic interference. the fiber optic sensor is essentially an optical fiber equipped with a reflecting grid (the so-called bragg grid), which is produced for example by uv laser firing during fiber production. it is a very small, germanium doped portion of the optical fiber, which in the case of joining this fiber with the stretched body changes its optical properties in a defined manner, and thus it is possible to measure very precisely the elongation of the body. on one fiber can be up to a few tens of independent fbg sensors, each measuring a different place on construction and working at a different wavelength. the fiber is equipped with a cover polymeric sleeve which acts as a fiber protection against moisture and mechanical damage. the fbg sensors are based on a principle of reflection central wavelength of light by the bragg grid. a part of a supplied signal passes through the grid and a part of the light spectrum is reflected back. in the measuring unit, it is compared to the reference (unloaded) grid tuned to the same wavelength, the deformation is obtained from the difference of values. in the case of a newly made glued laminated timber element, the sensor can be inserted into the wooden structure when it is glued. if an existing or even historical structure is to be monitored, the sensor can be attached to its surface. the system was tested under laboratory conditions when the fiber with the sensor was embedded into a beam structure of glued laminated timber. the beam was mechanically evaluated by a 4-point loading test until failure, see fig. 2a). the correct function of the sensor was examined and a obtained data were compared with an analytical approach, see fig. 2b). more detailed information about experiments and achieved results can be found in [6]. absolute error of measurement compared to analytical calculation is to 10%. from the comparison of the measured and calculated data it can be stated that the fbg sensors are also suitable for application to timber elements, for example for monitoring the stress of timber-concrete bridge structures. the system can be configured to track the behaviour of the load bearing element during the lifetime of the structure and if necessary, alert to structural overload or technical problems. the technology of fiber optic sensors could be one of the ways to control the state of the bridges in the future. 4. protection of timber by photocatalytic materials generally timber protection materials should prolong lifetime of timber elements. the main role of the group protective materials based on photocatalysis is to protect timber element before color changing, moulds and fungus. the most spread photocatalytic materials for wood protection are tio2, zno, ceo2 [7] and wo3. they are used in nano-form and usually applied as a transparent coating. there are two ways, particle can be directly mixed with the solution or encapsulated into porous, resistant material (e.g. silicates) and then apply on the timber. encapsulation 58 vol. 22/2019 monitoring and protection of timber-concrete bridges (a). (b). figure 2. laboratory testing of fbg sensors a) experimental arrangement of the test; b) comparison of measured and calculated data. should avoid damage of timber and save ability to decompose organic compound. nowadays is tendency made more efficient photocatalytic materials by dopping. dopping should caused decreasing band gap between valence band and conductive band which means that absorption expand in the visible region of light. until now the results in this research area are not satisfactory. dopped photocatalytic material are not such efficient like un-dopped one. less known is using of iron oxides. they are used in cases when it is not necessary keep wood color. other inorganic compound which were studied as a photocatalyzer were cds, cuo, moo3, cr2o3, but they are not used in bid scale. tio2 belong to the most known materials with the photocatalytic properties. it is mainly designed for decomposition of organic compound. a lot of materials with the photocatalytic properties were used earlier as a pigments or part of compound created pigments. it was observed that some of them, react with surroundings. exterior paints based on titania had tend to chalking which means that in surface stayed non-adhesive powder. tio2 is present in lot of minerals which were known already in ancient times but tio2 itself as a compound is known from nineteenth century. on the beginning of the twentieth century start industrial production. timber protection with tio2. tio2 was tested in different type of wood species different climatic conditions. the use of tio2 usually increased resistance against the fungi and mould. on the other hand it was mentioned in some works the use of tio2 negatively influenced mechanical properties. the morphology of tio2 is varied from spheres particle to tubes. in our focus is planar particles-leaflets of tio2. the type what we use is possible prepare as a photoactive and also as a non-photoactive. the synthesis of tio2 is simple but time consumed. the starting compounds are titanium sulphate nhydrate which is dissolved in deionized water. next step is precipitation by concentrated aqueous nh3. white precipitate is filtered, properly washed and transferred into a beaker and re-suspended in deionized water. after that is added h2o2 until solution became yellow. last step is lyophilisation. more details about synthesis is refer in literature [8]. we are also looking for the ways how to make it shorter and keep the morphology and same physico-chemical properties. the goal of our work is study interaction between planar particles of tio2 and wood matter. testing is running on the pine wood and beech wood. the samples were coated with different coating and concentrations of tio2. experimental data from uv-vis spectroscopy clearly show that the planar particle of tio2 absorbed wide range of uv light (uva, uvb) (fig. 3). tio2 was annealed on 200 °c, 250 °c and 300 °c due to remove residual peroxide after the lyophilisation. also we want to know whether the annealing temperature has an impact on the absorption of uv-light. tio2 annealed on 200 °c and 300 °c have the maximum on the same wave length (310 nm), maximum tio2 annealed on 250 °c is slightly shifted (320 nm). shape and maximum of the absorbance’s curves is in all cases similar. on the base of experimental data we assumed that in temperature range between 200 °c and 300 °c absorbance is same. the planar morphology was proved by electron microscopy (fig. 4). the surface of the particle is not plain but wavy and also it is not completely continuous. the holes cover 20% of the surface and aggregates 10%. holes and aggregates in this amount have no negative impact on the uv absorption of the material as was proven by uv-vis spectroscopy. 5. summary • timber-concrete composite bridges are a technical solution with great potential for bridge construction. the wide range of examples also prove that this 59 p. kuklík, l. velebil, a. gregorová, p. svora acta polytechnica ctu proceedings figure 3. uv-vis of planar, amorphous tio2. figure 4. uhrsem planar particle of amorphous tio2. technology very useful in the past but also has enormous potential for use in the future. • several aspects for future application of nondestructive techniques in bridge survey and restoration shall be considered. this may include enhancement of non-destructive methods accounting for specific applications in the structural assessment of existing bridges and preparation of draft proposals for technical guidelines in the field of bridges inspection. • the idea protect external timber construction with the photocatalytic material is not new. what is missing are information’s about the interaction between photocatalytic material and timber. we chose as a model material planar tio2 because it is possible prepare it in two forms amorphous (non-photoactive) and crystalline (photoactive). we can study interaction of two phases morphologically identical material which is very important because research community admit theory where the morphology is one of the main factor with big impact on the photocatalytic properties. we are able also study role of the coating solvent (water, water glass, acryl). in case of water solution we assume that water evaporate and the particles are bonded by van der waals force so we will study directly interaction particle – wood, in case of water glass solution the inorganic porous structure is created and particles are surrounded silicate chains (system to study particle silicate chain – wood) and in case of acrylic solution the organic porous structure is created (system to study particle – acrylic acid compounds). obtain data will be used for new evaluation method of influence photocatalytic materials on surface of timber structures, mainly timber bridges. acknowledgements this work is supported by the project ga18-26297s “study of interactions in system: wooden surface amorphus layer of tio2 crystalline layer tio2”. references [1] j. p. wacker, a. dias, t. k. hosteng. investigation of early timber–concrete composite bridges in the united 60 vol. 22/2019 monitoring and protection of timber-concrete bridges states. in proceedings, ictb 2017, 3rd international conference on timber bridges, p. 11. skellefteå, sweden, 2017. [2] h. hasníková, p. kuklík. various non-destructive methods for investigation of timber members from a historical structure. wood research 59(03):411–420, 2014. [3] b. kasal, r. anthony. advances in in situ evaluation of timber structures. progress in structural engineering and materials 6:94 – 103, 2004. doi:10.1002/pse.170. [4] a. kuklíková. kompozitní dřevobetonové konstrukce. ph.d. thesis, czech technical university in prague, faculty of civil engineering, prague, 2004. in czech. [5] j. maddox, p. kuklík. non-destructive testing methods for determining material characteristis values of historic timber structures. in m. c. forde (ed.), structural faults & repair 2016. 16th international conference and exhibition. edinburgh, 17.05.2016 19.05.2016. engineering technics press edinburgh, 2016. [6] l. velebil, r. zelený, j. včelák, et al. optical fibre sensors as a potential solution for monitoring wooden structures. in wcte 2016 cd-rom proceedings. vienna, 2016. [7] m. nikolic, j. lawther, a. sanadi. use of nanofillers in wood coatings: a scientific review. journal of coatings technology and research 12:445–461, 2015. doi:10.1007/s11998-015-9659-2. [8] j. šubrt, p. pulišová, j. boháček, et al. highly photoactive 2d titanium dioxide nanostructures prepared from lyophilized aqueous colloids of peroxopolytitanic acid. materials research bulletin 49:405 – 412, 2014. doi:10.1016/j.materresbull.2013.09.028. 61 http://dx.doi.org/10.1002/pse.170 http://dx.doi.org/10.1007/s11998-015-9659-2 http://dx.doi.org/10.1016/j.materresbull.2013.09.028 acta polytechnica ctu proceedings 22:57–61, 2019 1 introduction 2 evaluation of density and mechanical properties of structural timber 3 mechanical stress monitoring of timber-concrete bridges by fbg sensors 4 protection of timber by photocatalytic materials 5 summary acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0012 acta polytechnica ctu proceedings 22:12–16, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app short-term corrosion response of metals in lime mortar pavla bauerováa, b, zdeněk prošeka, c, jitka krejsováb, martin kepperta, b, ∗ a czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic b czech technical university in prague, faculty of civil engineering, department of materials engineering and chemistry, thákurova 7, 166 29 praha 6, czech republic c czech technical university in prague, faculty of civil engineering, department of structural mechanics, thákurova 7, 166 29 praha 6, czech republic ∗ corresponding author: martin.keppert@fsv.cvut.cz abstract. even though the corrosion of steel in concrete is very well described, the corrosion processes in lime environment studied very rarely. this environment is obviously also basic and its ph decreases in time, but the obtained results indicate important differences in behaviour of these two systems. corrosion response of carbon steel, zinc platted steel, copper, brass and lead placed in lime mortar was studied by help of sem microscopy. keywords: metal corrosion, modern mosaics, lime mortar. 1. introduction the problems related to the steel reinforcement corrosion in concrete are of crucial importance with respect to lifetime of majority of contemporary concrete structures [1]. there the most important corrosion-related feature is high ph, which, together with low potential, ensures formation of a passive layer of fe3o4 or fe2o3 on steel. the ph of concrete pore solution is controlled by soluble species present in concrete – portlandite (ca(oh)2) and eventually free alkalis (naoh, koh) [2]. in case that portlandite is the controlling species, the ph of pore solution is 12.5, corresponding to solubility constant of portlandite; the presence of alkalis can increase ph to higher level (ca. 13). obviously, also other components of pore solution are of importance towards the steel corrosion – especially chloride ions are widely studied [3]. the present paper is inspired by demand on knowledge of corrosion behavior of metallic materials in lime mortar; the demand came from restores dealing with mosaics. modern mosaics (from 19th and 20th century) have been frequently created in atelier by help of a metal frame, mortar and glass or stone tesseraes, and after the hardening were installed [4]. the metal frames, as well as other mosaics’ components, are frequently subjected to exterior conditions inducing complex degradation processes. the goal of the research was to evaluate corrosion behavior of carbon steel, copper, brass and lead in lime mortar environment. it is similar to the concrete environment – the ph is obviously also controlled by portlandite, but higher permeability of lime systems, responsible for their faster carbonation, makes the conditions somewhat different. corrosion tests are commonly performed by help of an accelerated test [5]; here just a short term response of metals to the lime environment was studied. 2. experimental the short-term corrosion behavior of five metals (carbon steel, copper, brass, lead) was studied. the strips (10 × 50 mm) of metal plates were degreased, rinsed and immersed in fresh lime mortar (lime hydrate + standard quartz sand 1:3) and left in laboratory conditions. the used lime hydrate was compliant with cl90s type; the xrf analysis revealed that cao content was 99.3%, mgo 0.5% and so3 0.1%. after three months, the metal strips were removed from the mortar, cross-cutted, fixed by epoxy resin, polished and the interface between metal and adhering mortar was examined by electron microscopy (sem phenom xl) in bse mode. the mortar was not yet fully carbonated; the xrd rietveld analysis (fig. 1) revealed that binder (aggregates excluded) contained 38% portlandite and 62% of calcite. it means that ph was still controlled by ca(oh)2 dissolution. 3. results and discussion carbon steel. the surface of carbon steel (fig. 2) has been covered by continuous layer of fe corrosion products fexoy; generally fe2o3 and fe3o4 may be formed at basic conditions, depending on the potential. the layer thickness ranged from 5 to 10 µm. the gap between metal and oxidized layer appeared during the sample preparation. mortar adhered well fe-oxides layer. the profile “1 fe” (line scan 1) shows that fe atoms has diffused also to the mortar layer. in areas, where mortar did not cover the steel well, a porous 12 http://dx.doi.org/10.14311/app.2019.22.0012 http://ojs.cvut.cz/ojs/index.php/app vol. 22/2019 short-term corrosion response of metals in lime mortar figure 1. phase composition of lime mortar under study. figure 2. interface between carbon steel and lime mortar. the arrows indicate line scans 1 and 2. dendritic fexoy structure appeared. the line scan 2 illustrates that the dendritic corrosion structure has grown also on the mortar surface. taking into account the short time (3 months) of experiment and laboratory conditions, one can state that steel corroded more that would be expected for ca(oh)2 rich environment. this behavior indicates, that in lime mortar is likely higher potential than in concrete; moreover in lime are not any alkalis like in cement which are responsible for higher ph than that corresponding to portlandite [6]. the zinc coated carbon steel was also subject of investigation, but one did not observed any corrosion effects. copper. the surface of copper (fig. 3) has been covered by well adhering layer of mortar; one observes that mortar is more compact in the vicinity of copper sheet. a closer view (fig. 4) indicates that the more compact surface layer is not homogenous; there are two phases. both of them are dominated by caco3, as results from ca and c profiles, but the phase in vicinity of copper plate is enriched by sulfur; its concentration is equal across the layer. the second element enriched in this layer is copper; its content is decreasing gradually in the layer and drops to zero on the interface between the two phases. the only source of sulfur in the system is lime, containing 0.1% of so3. due to low content, xrd diffraction was not able to detect, whether the is a sulfide or a sulfate salt. generally copper is immune in basic environment, or eventually, at higher potential, passivated by cu2o or cuo. in the studied system, copper surface is not seemingly corroded. on the other hand, certain amount of copper dissolved and formed a cu-s compound accumulating at the interface between copper and lime mortar. there has to be a driving force for diffusion of sulfur from bulk of lime to the interface; this driving force is a difference between sulfur activity 13 p. bauerová, z. prošek, j. krejsová, m. keppert acta polytechnica ctu proceedings figure 3. interface of copper (left) and lead (right) with lime mortar. figure 4. detail of copper – lime interface. the arrow indicate line scan. figure 5. elementary mapping performed on the copper – lime interface. 14 vol. 22/2019 short-term corrosion response of metals in lime mortar figure 6. sem micrograph of brass – lime interface. the arrow indicate line scan. figure 7. sem micrograph of lime mortar on lead substrate. the arrow indicate pb line scan. in bulk and at the interface. it indicates that a cu-s compound precipitated in the adjacent layer, what lowered the interface activity of a dissolved s-species, consequently inducing the s transport to the copper vicinity. this process can be observed also by elementary mapping of the adjacent cu-s enriched layer (fig. 5). brass. the interface between brass and lime mortar did not evince any indication of ongoing corrosion process (fig. 6). neither the element profiles did not indicate transfer of cu or zn from metal to the mortar phase. the mortar adhered well to the substrate. the results indicate that in the lime environment, zinc provides a passive layer to the brass and prevent its oxidation [7], even though in brass alloys, zn is usually corroding preferentially to copper due to its lower standard electrochemical potential [8]. lead. the adherence of lime mortar to lead was very poor (fig. 3). the distinct crystals in lead are sic particles, introduced, due to low hardness of lead, during the sample preparation. according to the lead pourbaix diagram, it dissolves at ph above 11.5 to pb(oh)2−4 anion [9]; correspondingly the lead gradient was detected in the mortar (fig. 7) where the formation of capb(oh)4 may be expected. 4. conclusions there is just a little attention paid, compared to concrete, to corrosion behavior of metals in contact with lime mortar. this environment is obviously also basic and its ph decreases in time, but the obtained results indicate important differences in behavior of these systems. carbon steel, which is passive in concrete, corroded in lime mortar already after 3 months; the corroded layer developed between metal and adhering mortar. the zinc coated steel did not corroded in lime environment within the studied time range. copper formed a cu-s corrosion product, dispersed in adjacent layer of mortar, despite it is relatively noble metal. on the other hand, zinc in brass provided 15 p. bauerová, z. prošek, j. krejsová, m. keppert acta polytechnica ctu proceedings passivation to the material and any corrosion has not been observed. the lead corroded to pb(oh)2−4 anion dispersed in lime mortar. it can be concluded that the best corrosion compatibility with the lime mortar environment proved brass and zinc coated steel. acknowledgements the work has been supported by czech science foundation project nr. 18-13525s “modern mosaic mortars in a microscope – methods for their materials characterization and degradation studies”. references [1] a. poursaee (ed.). corrosion of steel in concrete structures. elsevier ltd., 2016. doi:10.1016/c2014-0-01384-6. [2] a. behnood, k. van tittelboom, n. de belie. methods for measuring ph in concrete: a review. construction and building materials 105:176–188, 2016. doi:10.1016/j.conbuildmat.2015.12.032. [3] j. němeček, t. koudelka, t. krejci. numerical modeling of accelerated chloride transport in concrete. vol. 1978. 2018. doi:10.1063/1.5043737. [4] i. perná, t. hanzlíček, m. kracík štorkánová. characterization of historic mosaic at pfeiffer-kral sepulcher, jablonec nad nisou: a study of the mortar and tesserae origin. ceramics silikaty 58:308–313, 2014. [5] y. ma, z. guo, l. wang, j. zhang. experimental investigation of corrosion effect on bond behavior between reinforcing bar and concrete. construction and building materials 152:240–249, 2017. doi:10.1016/j.conbuildmat.2017.06.169. [6] a. poursaee. corrosion of steel bars in saturated ca(oh)2 and concrete pore solution. concrete research letters 1(3):90–97, 2010. [7] s. thomas, n. birbilis, m. venkatraman, i. cole. corrosion of zinc as a function of ph. corrosion 68:1–9, 2012. doi:10.5006/1.3676630. [8] s. selvaraj, s. ponmariappan, m. natesan, n. palaniswamy. dezincification of brass and its control an overview. corrosion reviews 21:41–74, 2003. doi:10.1515/corrrev.2003.21.1.41. [9] y. xie. dissolution, formation, and transformation of the lead corrosion product pbo2: rates and mechanisms of reactions that control lead release in drinking water distribution systems. ph.d. thesis, washingthon university in st. louis, 2010. 16 http://dx.doi.org/10.1016/c2014-0-01384-6 http://dx.doi.org/10.1016/j.conbuildmat.2015.12.032 http://dx.doi.org/10.1063/1.5043737 http://dx.doi.org/10.1016/j.conbuildmat.2017.06.169 http://dx.doi.org/10.5006/1.3676630 http://dx.doi.org/10.1515/corrrev.2003.21.1.41 acta polytechnica ctu proceedings 22:12–16, 2019 1 introduction 2 experimental 3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0125 acta polytechnica ctu proceedings 13:125–129, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app hydration process and mechanical properties of cement paste with recycled concrete powder and silica sand powder jaroslav topiča, ∗, zdeněk prošeka, b a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b university centre for energy efficient buildings, czech technical university in prague, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: jaroslav.topic@fsv.cvut.cz abstract. recycled concrete powder (rcp) mostly consisting of cement paste could be reused as partial cement replacement. the aim of this paper is to compare hydration and mechanical properties of rcp and two types of silica sand powder (ssp). comparison of those materials combined with cement can highlight the binder properties of recycled concrete powder. using of two types of ssp also show an influence of their fines on hydration process and mechanical properties. particle size analysis and calorimetric measurement were carried out and mechanical properties such as bulk density, dynamic young’s modulus and compression strength were examine. calorimetric measurement proves the presence of exposed non-hydrated particles in rcp that can react again. however lower density of old cement paste in rcp overweight the mentioned potential of rcp and mechanical properties are decreasing compared with reference cement paste and cement paste ssp. keywords: cement paste, recycled concrete powder, silica sand powder, hydration process, mechanical properties. 1. introduction the great popularity of cement composites such as concrete in the building industry, among other things, also bring difficulties of recycling old materials from demolitions and reconstructions. recycling process of old concrete reach a good effectivity of reusing of it. old concrete is usually used in earth works for backfill, sub-base layers of the roads [1, 2], unpaved road surfaces [3, 4] or concrete production [5–7]. using old concrete such way result to amount of not usable powder. this powder mostly consist of cement paste which could be reused for remaining binding properties of non-hydrated core of cement grains. in previous papers where the recycled concrete powder (rcp) were used as partial cement replacement and micro-filler [8, 9]. the aim of this paper is to compare hydration and mechanical properties of rcp with two inert fillers with similar range of granulometric curve or similar specific surface area. to eliminate influence of larger nucleation area (caused by presence of fine particles in rcp) on hydration process the inert silica sand powder was used to comparison. due to his crystalline character, the pozzolanic activity of sio2 should be minimal compared with amorphous sio2, which is used as reactive additive [10, 11]. comparison of those materials combined with cement can highlight the binder properties of rcp and influence of recycled concrete powder on hydration process can be defined [12, 13]. using of two types of silica sand powder also show an influence of their fines on hydration process and mechanical properties. results of those two types silica sand powder should help to further modification properties of recycled concrete powder. 2. materials and samples all mixtures for sample production composed of the portland cement cem i 42.5 (radotín, czech republic) and different amount of following admixture. samples detonate cem+rcp contained recycled concrete powder (rcp) produced by crushing and further micro-milling of old concrete railway sleepers. samples detonate cem+ssp contained silica sand powder st2 (ssp2) or st6 (ssp2) from company sklopísek střeleč, inc. four different mixtures were used for calorimetric measurement. first mixture (denote as cem) was reference cement past. second mixture (denote as cem+rcp 50) contained 50 % of cement and 50 % of rcp. third mixture (denote as cem+ssp2 50) contained 50 % of cement and 50 % of ssp st2 and fourth mixture (denote as cem+ssp6 50) contained 50 % of cement and 50 % of ssp st6. all samples had same water/mixture ratio equal to 0.45. for this measurement approximately 35 g of the mixture was used and put into plastic container. seven different mixtures were used for evaluation of mechanical properties. composition of the mixtures is shown in table 1. mixtures had different water/mixture ratio from 0.35 to 0.40 to compensate the 125 http://dx.doi.org/10.14311/app.2017.13.0125 http://ojs.cvut.cz/ojs/index.php/app jaroslav topič, zdeněk prošek acta polytechnica ctu proceedings mixture cement [g] recycled concrete silica sand water/mixture (cem i 42.5 r) powder (rcp) [g] powder (ssp) [g] ratio cem (ref) 1500 0.35 cem+rcp 10 1350 150 0.36 cem+rcp 50 750 750 0.40 cem+ssp2 10 1350 150 0.36 cem+ssp2 50 750 750 0.40 cem+ssp6 10 1350 150 0.36 cem+ssp6 50 750 750 0.40 table 1. composition of the mixture. figure 1. the particle size distribution of rcp, ssp st2 and ssp st6. deteriorated workability caused by increasing amount of rcp and ssp. each set contained 6 prismatic samples having dimensions of 40×40×160 mm. the samples were removed from their casts after 2 days. after that the samples were cured for 28 days in water at the temperature of 21 ± 2 ◦c. 3. experimental methods rcp particle size distribution and determination of middle grain size were performed using a fritsch analysette 22 microtec plus laser device. a matest e009 device was used to measure the specific surface area using the blain method. calorimetric measurement was conducted on an isothermal tam air calorimeter for measurement of heat flow. mixtures were tested for 7 days at a constant temperature of 20 ◦c and were stored in sealable plastic containers, each containing from 32 g to 36 g of mixture. based on the weight of each sample (measured before testing) the results of the heat flow were related to 1 g of cement for a better detection of changes caused by admixtures. bulk density of the samples was evaluated from dimension and weight of the samples before resonance testing. dimension and weight of the samples was used as input values for resonance testing. dynamic young’s modulus was monitored by using the non-destructive resonance method. this method is based on measuring the natural frequency of the prismatic 40×40×160 mm samples. for evaluation of dynamic young’s modulus the basic longitudinal and flexural natural frequency, dimension and weight of the samples were used. compressive strength was determined on 28 day old samples using a model fp100 heckert device. the testing was displacement controlled at a constant rate of 0.1 mm/s in the case of three-point bending and 0.3 mm/s during the compression test. a uniaxial compressive test was performed on the broken specimen halves with effective dimensions 40×40× ∼80 mm 4. result and discussion in figure 1 are presented results of granulometric measurement. as can be seen rcp has high amount of very fine particles but also has long range. ssp st2 was chose for comparison with rcp because of his high range of granulometric curve. ssp st6 was chose for comparison with rcp because of his similar fines. specific surface area of cem i 42.5r was equal to 10811 cm2/cm3 and very similar to rcp with value 10996 cm2/cm3. ssp st2 has specific surface area equal to 8676 cm2/cm3 and ssp st6 to 11604 cm2/cm3. results of calorimetric measurement are 126 vol. 13/2017 hydration process and mechanical properties ... figure 2. heat flow of cement paste and cement with 50 wt. % of recycled concrete powder (cem+rcp 50) and silica sand powder (cem+ssp2 50 and cem+ssp6 50). samples bulk density samples bulk density [kg/m3] [kg/m3] cem (reference) 1919 ± 9 cem+rcp 10 1851 ± 1 cem+rcp 50 1566 ± 3 cem+ssp2 10 2026 ± 11 cem+ssp2 50 1964 ± 3 cem+ssp6 10 1906 ± 6 cem+ssp6 50 1722 ± 5 table 2. bulk density of the samples. figure 3. dynamic young’s modulus of cement paste and cement paste with 10 and 50 wt. % of recycled concrete powder (cem+rcp) and silica sand powder st2 and st6 (cem+ssp2 and cem+ssp6) after 28 days. 127 jaroslav topič, zdeněk prošek acta polytechnica ctu proceedings figure 4. compression strength of cement paste and cement paste with 10 and 50 wt. % of recycled concrete powder (cem+rcp) and silica sand powder st2 and st6 (cem+ssp2 and cem+ssp6) after 28 days. shown in figure 2. by relating heat flow to 1 g of cement and not an entire rcp or ssp mixture it was possible to highlight differences in the hydration process caused by rcp. in case of mixtures cem+ssp2 50 and cem+ssp6 50 the increase of heat flow compared with reference cement was caused by presence of ssp which provided larger nucleation area to cement hydration. cem+ssp6 50 had little bit higher increase of heat flow compared with cem+ssp2 50 thanks to his higher specific surface. increase of heat flow of mixture cem+rcp 50 was partly (as in cem+ssp) caused by larger nucleation area, but mostly by presents of non-hydrated cement grains in rcp. increase of heat flow of mixture with rcp compared with mixture cem proves that exposed cement grains in rcp can react again and comparison with cem+mss 50 proves that increase of heat flow wasn’t caused only by larger nucleation area. results of bulk density of the samples are shown in table 2. there is obvious decrease of bulk density of the samples with rcp in both cases compared with reference samples or samples with. the high amount of fine particles and lower density of old cement paste compare with the new one caused decrease of bulk density in case of cem+rcp. samples cem+ssp6 10 exhibit comparable bulk density and cem+ssp6 50 had lower bulk density then reference samples. higher fines also cause the decrease of bulk density in case of samples with ssp st6 but decrease was not so steam because of higher density of ssp compared with rcp. addition of coarser ssp st2 resulted to increase of the bulk density in both concentrations. dynamic young’s modulus of the samples illustrated in figure 3 and compression strength illustrated in figure 4 closely related to bulk density. as can be seen the mechanical properties of samples with rcp are despite the hydration of cement grains in rcp decreased compared with other samples. this decrease was probably caused by lower density of old cement paste in rcp compared with ssp and new hydration products and simultaneously by high amount of very fine particles. dynamic young’s modulus and compressive strength of the samples cem+ssp2 10 increased compared with cement paste due to increased cohesiveness of matrix caused by grains that effectively work as micro filler. based on those two mechanical properties the ssp st6 in samples cem+ssp6 10 has similar micro filler effect but the specific surface area of was too high and contain higher amount of very fine particles that cause small decrease of compression strength. in all samples with 50 wt % of rcp or ssp there was huge decrease of compression strength cause by inappropriate ration of cement and filler. based on results, the potential improvements of mechanical properties could be achieved by using small amount of rcp in cement mortars or concrete. in this combination, the rcp would be fulfills the role of micro filler and partially the role of binder and lower density of rcp would not be obstacle as in case of cement paste. thanks to using two types of ssp with different fines can be assumed that better results would be achieved by using coarser rcp with similar fines as ssp st2. remaining question is if non-hydrated cement grains in old cement paste can exposed on this level and still fulfills the role of partial binder replacement. experiment that answers this question will follows in further research. 5. conclusions this paper is focused on influence of different type of additive and its properties on hydration and mechanical properties of cement paste. in this experiment the rcp was used to fulfills the role of micro filler and partial cement replacement. also two types ssp was used that thanks to its crystalline character works as inert micro filler. based on results the following findings are summarized: • based on heat flow during hydration process, the presence of exposed non-hydrated cement grains 128 vol. 13/2017 hydration process and mechanical properties ... in rcp has been proved and influence of larger nucleation area was also define as minimal thanks to comparison with influence of ssp with different fines. • despite the additional hydration of cement grains in rcp the mechanical properties of cement paste with rcp were negatively influenced by lower density of cement paste in rcp compared with ssp and new hydration products. • based on result two types of ssp the negative influence of high amount of very fine particles was recognized. • the rcp can be used as micro filler and binder replacement only in small amount into cement composites where aggregate is present. this way the influence of lower density of old cement paste can be eliminated. • based on results of samples with ssp can be stated that using rcp with lower amount of very fine particles and lower specific surface area might lead to improvement of mechanical properties of cement paste. but binder properties can be diminished by insufficient exposure of non-hydrated cement particles. acknowledgements the authors wish to express their gratitude for financial support to the faculty of civil engineering, czech technical university (ctu) in prague under sgs project no. sgs16/201/ohk1/3t/11 and no. sgs17/060/ohk1/1t/1 and czech science foundation gačr under project no. 17-06771s. the authors also thanks to the center for nanotechnology in civil engineering at the faculty of civil engineering, ctu in prague. references [1] a. barbudo, f. agrela, j. ayuso, et al. statistical analysis of recycled aggregates derived from different sources for sub-base applications. construction and building materials 28(1):129–138, 2012. [2] c. j. engelsen, g. wibetoe, h. a. van der sloot, et al. field site leaching from recycled concrete aggregates applied as sub-base material in road construction. science of the total environment 427:86–97, 2012. [3] i. vegas, j. ibañez, a. lisbona, et al. pre-normative research on the use of mixed recycled aggregates in unbound road sections. construction and building materials 25(5):2674–2682, 2011. [4] j. r. jimenez, j. ayuso, f. agrela, et al. utilisation of unbound recycled aggregates from selected cdw in unpaved rural roads. resources, conservation and recycling 58:88–97, 2012. [5] r. silva, j. de brito, r. dhir. properties and composition of recycled aggregates from construction and demolition waste suitable for concrete production. construction and building materials 65:201–217, 2014. [6] f. agrela, m. s. de juan, j. ayuso, et al. limiting properties in the characterisation of mixed recycled aggregates for use in the manufacture of concrete. construction and building materials 25(10):3950–3955, 2011. [7] l. evangelista, j. de brito. mechanical behaviour of concrete made with fine recycled concrete aggregates. cement and concrete composites 29(5):397–401, 2007. [8] j. topič, z. prošek. properties and microstructure of cement paste including recycled concrete powder. acta polytechnica 57(1):49–57, 2017. [9] j. topič, z. prošek, t. plachý. influence of increasing amount of recycled concrete powder on mechanical properties of cement paste . [10] p. barnes, j. bensted. structure and performance of cements. crc press, 2002. [11] s. shebl, h. seddeq, h. aglan. effect of micro-silica loading on the mechanical and acoustic properties of cement pastes. construction and building materials 25(10):3903–3908, 2011. [12] s. guettala, b. mezghiche. compressive strength and hydration with age of cement pastes containing dune sand powder. construction and building materials 25(3):1263–1269, 2011. [13] j. šeputytė-jucikė, m. kligys. the effects of modifying additives and chemical admixtures on the properties of porous fresh and hardened cement paste. construction and building materials 127:679–691, 2016. 129 acta polytechnica ctu proceedings 13:126–130, 2017 1 introduction 2 materials and samples 3 experimental methods 4 result and discussion 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0031 acta polytechnica ctu proceedings 11:31–34, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app railway traffic optimization dušan kamenický∗, tomáš brandejský, vít fábera, valerii gopak faculty of transportation sciences, czech technical university in prague, konviktská 20, 110 00 praha 1 ∗ corresponding author: kamendus@fd.cvut.cz abstract. the article focuses on analysis of possibilities of operatic control of railway traffic and its automatization and its support by information systems. conditions, necessary pieces of input information like infrastructure description are discussed. railml format is mentioned as suitable format for description in railway area. optimization criterion and use of evolutionary techniques are suggested. the system is modeled in railway laboratory of the faculty of transportation sciences, ctu in prague. keywords: railway traffic optimization, railml, evolutionary techniques, prediction. 1. railway traffic optimization trends increasing demands to railway capacity require to build new railway lines and to apply new approaches like information technologies supporting railway traffic to use effectively given infrastructure. train running is controlled and checked by interlocking systems (and by operations over the devices) that evaluates the state of the infrastructure (mainly occupancy of track sections is detected using track circuits or axle counters), controls infrastructure elements (switches), transmits instructions how to move through signals or automatic train stopping device. there are several trends of rail operation control. firstly, the way of control is shifted from distributed system spread over stations to centralized one in all europe. the most railway lines in the czech republic should be controlled from two offices in přerov and prague. the centralization of the operation control gives spread view of the situation and more information to the operator. on the other hand, there is a disadvantage of this approach. information quantity to evaluate increases if the area is large. it is clear the information capacity of each human operator is limited and different. moreover, operator is negatively influenced by rotes. secondly, next modern trend (especially in switzerland, belgium and germany) is a separation of strategy control (route planning from daily to annual interval) from operative control (conflicts solution caused by deviations from regular time-table due to infrastructure or vehicle faults) and from direct control (safety device operation, shunting). each level of operation control architecture process specific functions and it needs it support. this article is focused especially on operative control and its it support, strategy control is mentioned only marginally. it is supposed that output of the strategy level support is a scheduled daily time table with valid and conflict-free train routes (really realizable with respect to the route capacity). operative control it support should react to deviation from planed daily time table; it monitors actual trains routes and solves potential conflicts with others trains in advance (from minutes to hours). the output is modified conflictfree time table – specific route plan transmitted to direct control level (realized by human operator or automatic system). the direct control level is freed from searching the sequence of runs if the planned time-table is disturbed and processes of the direct control level are focused on safety device operations and guarantee of rail operation safety. 2. it support of operative control it support of operative control was missed out in the czech republic for a long time. today systems only shows planned timetable which is modified based on information about actual train position entered by operators in stations. they reflect neither train properties (parameters) nor infrastructure properties. it depends only on abilities and experiences of operators (dispatchers) if they are able to predict potential conflict and to suggest appropriate solution. moreover, todays systems are not able to transfer such modified timetable to the direct level operating on the plant-interlocking devices. to increase capacity of railway traffic and effectivity by optimization of operative control, it means to solve several partial problems that are very difficult: prediction of train paths, conflicts detection and optimal solution of these conflicts. the necessary premise to be able to search automatically optimal solution is to predict train path. some problems hamper to solve this problem satisfactory. above all, it is necessary to dispose enough description of the infrastructure. the description must contain the information related to the construction of time table (position and height of platform edges, line parameters affecting train resistance – slope, curves, tunnels, ...) and related to the interlocking system (track sections and their dependencies ensuring safety 31 http://dx.doi.org/10.14311/app.2017.11.0031 http://ojs.cvut.cz/ojs/index.php/app d. kamenický, t. brandejský, v. fábera, v. gopak acta polytechnica ctu proceedings running). the prediction is also affected by actual train parameters and their correctness, especially by type of motive power and its traction characteristic, train weight, braking ability of the train (both parameters can change in dependence of weather), eventually coefficients of train resistances. the strategy of run is important as well – ratio of times of thrust/coasting/braking. if the trains are manually controlled by engine drivers the strategies can be dramatically different. the application of ato (automatic train operation) brings the improvement in this area, when strategies became more stable (but the strategies can be different within several ato systems depending on producers). very stochastic quantity is time of stop in stations (getting on and off, operations on train). one possibility to eliminate partially randomness is to collect long-term statistics. dependencies between trains are necessary to take into account – connections, locomotives, coaches and train crew turnround cycles. completeness and up-to dateness of data stored in is are necessary. train routes conflict detection is highly dependent on the interlocking systems. each infrastructure operator has different equipment and it works with different set of risks and their consequences. due to this reason, each operator defines different functional requirements for interlocking systems and procedures to guarantee safety of traffic. different requirements can affect conditions for compatible routes. conditions depend not only on infrastructure but also on equipment of rolling stocks. it is not sufficient to consider only occupation of track sections but locking of sections which are running – before running and after running during defined interval. route locking can have time and space aspect. the specific case is such running conflict when one train run through the station or it is coming to the platform and the arrival to the second platform is at the same level (there is no subway or upperbridge). if a conflict is detected it must be solved. while previous phases (train path prediction and conflict detection) can be defined, modeled and computer relatively exactly, the definition of the optimization criterion by conflicts solution is complicated and "soft". one criterion can be "accuracy of planned time timetable", i.e. minimum of total sum of train delays. afterwards, weights of several trains and stations must be defined (delay of ic train has greater impact). but, weight assignment trenches offering politics of carriers. energetic criterion can be more objective and better definable. power consumption is possible to be good computed. the final criterion function can be multicriterial and be of the weight sum form. requirement to minimize time of closing level-crossing can be included to the criterion function. the set of situation is large that can happen. modern systems are able to solve only several situations automatically and sophistically. it is always considered a human operator and automatic system cooperation. own intervention can be realized not only by time shift of train path but also/or by change of train route – only within the station (to the different line) or the new whole path. in the case of new path there are new constraints as knowledge of line by driver, ability to pass new line by the train, operation of stations, ... 3. implementation in railway laboratory railway laboratory was founded at the faculty of transportation sciences which serves as a place to simulate train traffic control and interlocking system. the laboratory is equipped by interlocking devices typically used in the czech republic – from the oldest (mechanical) ones up to modern (computer based) ones. so, the level of direct control is modeled realistically. the laboratory is extended by higher levels traffic control architecture, namely on operative level. applications are developed which reflect modern trends of rail traffic control, focusing on train path prediction, conflicts detection and their solution. the topology of model railway in railway laboratory was selected as infrastructure on which applications will be developed and tested. we analyzed current concepts of infrastructure description used in reality. the way used in the czech republic prefers time-table point of view. this way is not compatible with the format representing infrastructure from the interlocking architecture point of view. the railml.org initiative [1] tries to uniform the infrastructure description and defines railml standard. the railml standard is based on xml format which is generally used nowadays to store structured data. it is a text format consisting of elements that can be nested; element is a block bounded by tags enclosed in <>. typically, configuration files, some database files have xml structure (html file are types of xml too). the version railml 2.3 was released, the specification railml 3 is in progress – the version 3.0 is released only for internal use in initiative, the version 3.1 is supposed to be released in autumn 2017. the railml defines several subschemas to describe and store information of various railway subsystems and parts: timetable and rostering, infrastructure, rollingstock, interlocking and common. a fragment of railml code 2.3 describing track is shown bellow (the complete example is available at [2]) – fig. 1. experiences show that extraction of infrastructure and connections from railml 2.x is not so clear especially due to separate description of elements and their connections. the railtopomodel project exists concurrently to the railml and it is a logical object model to standardize the representation of railway infrastructure-related data and it uses railml 3.x; the schema railml 3.x is different from 2.x and much better. nevertheless, version 3, as mentioned above, is still in progress. therefore research team decided to utilize existing own xml format describing 32 vol. 11/2017 railway traffic optimization figure 1. description of connection using railml 2.x (source [2]). figure 2. optimization system and its inputs. infrastructure which is used for interlocking devices configuration in railway laboratory [3]. optimization of the timetable is understood dynamic change of timetable (including time shift and route changes) to solve conflicts in our conception, not creation of the new (for example, daily, annual timetable) one. let’s summarize inputs of the optimization system (fig. 2): infrastructure description, train description, timetable description. infrastructure of railway laboratory is described by own xml format, as mentioned above. the net is composed from several types of elements like: tracks sections, signals, insulated rail joints, switches. the model contains other elements which are omitted in our application due to irrelevancy: power supplies, derailers, track circuits. elements and their bindings are read from the file and an oriented graph is constructed in the computer memory. own library was implemented in c++ programming language due to its speed and effective work with pointers. the base of own xml parser to read infrastructure description is an open-source library. there are many general freeware libraries for xml format analysis. we chose xerces library which is also written in c++ and it holds dom (document object model) in its internal structures. it is supposed that railml will be widely applied in real railway in the future. similar parser of railml format can be created and it is not a big problem to program such software module (library) "railml parser" containing functions for the analysis of the railml files – just due to existence of free general xml parsers like xerces. train description has our proprietary format. parameters of specific trains are stored to compute their traction and another physical characteristics (maximal speed) allowing modeling of train move and change it: increment or decrement train speed as one possible solution of collisions. naturally, the railml format concept calculates to store rolling-stock description in xml. timetable is stored in mysql server in railway laboratory (the future is railml format as well) and software library to read the timetable was created. the database contains only stations passed through and arrival and departure times. amount of information is insufficient. it is not possible to derive uniquely train path (there exist more paths between stations very often) and times of track sections occupancy. the identification of collision is not possible without knowledge of this piece of information. the time-table database has to be extended to contain sequence of track sections passing by trains and intervals of occupancy. each track section is labeled during simulation if it is free, occupied or locked. conflicts are searched like intersections of time intervals. if conflict is detected time intervals can be changed (shifted, shortcut, stretched) by the change of the train speed with using train characteristics. conflict solution can be implemented by some heuristic (evolutionary technique). if the change of travelling time doesn’t lead to the conflict solution an alternative paths can be used (which can generate unfortunately additional conflicts). alternative path searching can be implemented by evolutionary technique too. the technique of travelling time change can be applied to alternative paths. the conflict is solved by the path change of one train in the simplest case. if it is not solved the area of changes is spread (to other trains). it is evident the task is very complicated and it must be solved in reality by parallel computation. 4. conclusion the effective railway traffic optimization and following control is very difficult task. it is necessary to have enough described infrastructure, time table and rolling stock and store the data in information systems. prepared railml format and railtopomodel appears to be very perspective. to optimize railway traffic dynamically, especially when unexpected events happen, it needs to predict train runs, search conflict and solve them, for example to shift runs in timetable or to plan alternative paths. the optimal solution can be searched by evolutionary techniques where optimization criterion can be to minimize total sum of train delays or minimize power consumption. parallel computation is supposed to be applied. references [1] railml.org. initiative [online]. [cit. 2017-03-08] "http://www.railml.org". 33 d. kamenický, t. brandejský, v. fábera, v. gopak acta polytechnica ctu proceedings [2] railtopomodel [online]. [cit. 2017-03-08] "http://www.railtopomodel.org". [3] p. koutecký. electronic signal box for ctu transport lab, diploma thesis. faculty of electrical engineering, 2011. 34 acta polytechnica ctu proceedings 11:31–34, 2017 1 railway traffic optimization trends 2 it support of operative control 3 implementation in railway laboratory 4 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2017.10.0016 acta polytechnica ctu proceedings 10:16–33, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app testing methods for mechanically treated soils: reliability and validity ana petkovšek∗, matej maček, jasna smolar university of ljubljana, faculty of civil and geodetic engineering, jamova 2, ljubljana, slovenia ∗ corresponding author: ana.petkovsek@fgg.uni-lj.si abstract. a possibility of in-situ mechanical improvement for reducing the liquefaction potential of silty sands was investigated by using three different techniques: vibratory roller compaction, rapid impact compaction (ric) and soil mixing. material properties at all test sites were investigated before and after improvement with the laboratory and the in situ tests (cpt, sdmt, dpsh b, static and dynamic load plate test, geohydraulic tests). correlation between the results obtained by different test methods gave inconclusive answers. keywords: liquefaction, experimental site, rapid impact compaction, soil mixing. 1. introduction the term ground improvement refers to any technique or process that improves the engineering properties of the treated soil mass in order to fulfil the purpose intended. usually, the modified properties are strength, stiffness, permeability, workability (willingness for compaction), resistance to frost, etc. [1–4]. the type of ground improvement/soil treatment depends on the demands of the structure to be built, the soil type, the available materials and tools, and the local tradition and experience. the efficiency of a specific improvement technique may be measured by using different testings and instrumentation. whenever a laboratory or field test/measuring device is used as part of the data collection process, the validity and the reliability of the test/device is important. just as we would not use a test for the determination of particle size distribution to assess the activity of clay, we would not want to use any measuring device not capable of measuring properties which are of the design or qa – qc interest. this paper focuses on the validity and sensitivity of the laboratory and field tests used to assess the efficiency of the mechanical improvement of siltysandy soils. special attention is given to the use of kozeny-carman equation for the estimation of the soil permeability. analysis presented in the paper refers to the investigation performed at the location of hydropower plant (hpp) brežice, se slovenia, where the top layer of the foundation ground indicates properties of liquefiable soil. the main goal of the ground improvement is to increase the soil liquefaction resistance and to assure the required permeability of the fill material to be used for app. 14 km long dikes along the hpp reservoir. 2. background 2.1. site description the reservoir of the hpp brežice requires the construction of app. 7 km long and up to 10 m high dikes on both sides of the river sava. the top layer of the foundation ground consists of loose silty sands, prone to liquefaction. hpp brežice is the 5th in the chain of the lower sava river hpps (figure 1) and is situated in the influential area of nucler power plant krško (npp). tectonically, the wider area belongs to the tectonic unit named “krško basin”, which is considered a potentially active structure [5]. a well-documented earthquake with a magnitude of 5.7 and intensity of viii ems occurred in january 1917 [6]. the so called “zagreb earthquake” in 1880 triggered liquefaction along the river bank between brežice and zagreb (croatia) [7, 8]. the design ground accelerations in the krško basin for different earthquake return periods are presented in table 1. 2.2. geological profile the geological profile is quite uniform over the entire area of the hpp (figure 2). the top layer is up to 5 m thick, made of very loose sandy silts and sands (ml, sm, sp), deposited by recent floods. the top soil layer contains fine coal particles that origin from the flotation processes from the area around 60 km upstream where the coal mines have operated for the last 300 years. coal particles confirm the very young origin of the sediments. poplar tree roots appear up to 4 m deep and also prove the development of the “paleo” ground between the larger flood events. beneath the top layer the medium dense to dense quaternary gravels are deposited. the gravel is classified as silty sandy gravel (gm) to poorly graded gravel (gp) and consists mainly of limestone grains. 16 http://dx.doi.org/10.14311/app.2017.10.0016 http://ojs.cvut.cz/ojs/index.php/app vol. 10/2017 testing methods for mechanically treated soils table 1. design ground accelerations for different earthquake return periods [9–11] figure 1. location of the hpp brežice. the thickness of gravel deposits varies between 8 and 12 m. the pre-quaternary bedrock below the gravel consist of miocene sediments, represented by overconsolidated, not cemented silts and marls, with some inclusions of the weak, soft limestones. the ground water (gw) level is controlled by the sava river. during most of the year, the gw level appears app. 4 – 6 m below the ground surface, inside the gravel layer. during the floods, the gw may rise up to 2 m above the ground and may stay there for days. when the hpp brežice is put into operation, the existing ground surface and silty/sandy layers will be permanently flooded in the area between the dikes on the left and the right river bank. the permanent rise of the gw level is forecast also at the influential area from brežice to krško, on the dry sides of the dikes. 2.3. soil properties 2.3.1. index properties the grain size distribution of the quaternary soils is not uniform. in the top layer they may vary from sandy silt to poorly graded sand (figure 3). according to empirical data [12, 13], the top layer soils belong to the groups of most liquefiable and potentially liquefiable soils. index properties of typical soils from the top layer are given in table 2. laboratory proctor compaction tests show that due to the presence of coal particles, unfavourable grading and uneven water content distribution, the top layer soils willingness to compaction is poor. different compaction curves (figure 4) confirm heterogeneity of different strata, deposited inside the top layer. 2.3.2. liquefaction potential the spt testsperformed in top soil layer shown the modified blow count (n1)60 between 2 and 8. seed and idriss [14] proposed an empirical method for assessing 17 a. petkovšek, m. maček, j. smolar acta polytechnica ctu proceedings figure 2. schematic geological profile (left) and schematic cross section of the valley after the hpp construction (right). not in scale. figure 3. relationship between grading and liquefaction potential [13], with the boundaries of the grading of top layer and quaternary gravel at the hpp brežice. grading curves *nova loza and *moj dvor, reported by veinović et al. [7] belong to the area, where liquefaction appeared in 1880. figure 4. proctor curves for three typical groups of soils from the top layer. 18 vol. 10/2017 testing methods for mechanically treated soils table 2. properties of three typical groups of soils from the top layer. the likelihood of liquefaction of level sites (figure 5, left). the cyclic stress ratio (csr) is calculated from eq.1 [9]. (csr)m =7,5 = csr msf = β ( σv0 · amax σ′v0 ) · rd msf (1) msf = 6.9 exp (−m 4 ) − 0.058 msf ≤ 1.8 (2) where β is a ratio between real maximum and equivalent seismic stresses (0.65) [15], amax is the maximum ground surface acceleration on the foundation ground, rd is stress reduction factor that drops from 1 at the surface to about 0.9 at 10 m depth. msf is a factor to account for the varying number of cycles associated with different magnitude events and has values shown in table 1. the seismic dilatometer tests (sdmt) in top layers gave the shear wave velocities (vs) of 160-300 m/s, while the seismic refraction test indicates the vs of 90-110 m/s. seed and idris [14], youd et al. [16] and andrus [17] proposed an empirical relation for assessing the likelihood of liquefaction of level sites, based on relationship between the csr and vs (figure 5, right). the cyclic simple shear test (csst) was used to study the influence of the density on the liquefaction potential (figures 6,7). the tests were carried out on soil samples with properties given in table 2. specimens were prepared at different initial void ratios, saturated, consolidated at 75 kpa of effective stress and then loaded with horizontal sinusoidal load at desired stress ratio in undrained condition. the beginning of liquefaction was defined when the increase of pore water pressure was equal to 95 % of effective vertical stress before cyclic loading. the reference number of cycles was defined after kramer [18], for an earthquake of magnitude 7.5. results of the laboratory tests are given in figure 7 and confirm that the top layer soil are prone to liquefaction. the compacted soil samples at 95 % of maximum proctor density (dpr) will very probably resist the 475 year return period earthquakes and will very probably liquefy at stronger earthquakes. 2.4. preliminary conclusions the grain size distribution, the spt, the seismic refraction as well as csst indicate the same ranges of liquefaction potential and can be estimated as valid and reliable. the test results can also be confirmed by the historical data from the 1880’ zagreb earthquake. the sdmt test gave the shear wave velocities (vs) in a range of 160 – 300 m/s. comparing the sdmt test results with other tests as well as with calculation, using the empirical equation 3 and 4, explained in chapter 4, we may conclude, that the sdmt test values are too high for the top soil layer probably affected by field site heterogeneity or by the influence of other men activities. based on other tests the conclusions are clear: the top layer represents a very unfavourable foundation ground and must be improved or removed from the influential area of the dikes. 3. discussion on improvement technique several options were discussed to improve the liquefaction resistance of the foundation ground, as follows: • loose material from the top layer shall be removed and replaced with compacted gravel material. the idea was recognized as the most promising and reliable choice as large quantities of good quality gravel material are available at the location. however, due to the large volumes of top soil to be replaced (length of app. 14 km, depth up to 5 m, width that locally exceeded 50 m) the decision was made to investigate other options too. 19 a. petkovšek, m. maček, j. smolar acta polytechnica ctu proceedings figure 5. relationship between cyclic stress ratios causing liquefaction and (n1)60 (after seed and idriss [14]) and the position of the top soil in it (left). relationship between cyclic stress ratios causing liquefaction and shear wave velocity values (after seed and idris [14], youd et al. [16] and andrus [17]) and the position of the top soil in it (right). bc – before compaction, fc – fines content, ts – top soil layer. figure 6. the seiken cyclic simple shear apparatus, used for the laboratory tests (left) and the cell (right). figure 7. the relationship between the number of cycles needed for liquefaction and the cyclic stress ratio (csr) for top soils at different void ratios. 20 vol. 10/2017 testing methods for mechanically treated soils • option 1 (es 1): static and vibratory roller compaction. the main attention was paid to the following questions: (a) is the 95 % of max. proctor density attained on site, (b) what is the effective depth of the roller compaction and (c) what are the potentially harmful effects of coal particles and roots to the soil compaction behaviour? the four test fields were made to study the compaction efficiency of the foundation ground and up to 2.5 m high test embankment was constructed to study the roller compaction efficiency above the soft subgrade, the vertical and horizontal movements, the internal erosion and to analyse the sensitivity of different test methods, like cone penetration test (cpt), flat dilatometer test (dmt, sdmt), dynamic probing super heavy (dpsh b), etc. the test field was made in september and october 2009. • option 2 (es 2): rapid impact compaction (ric). to avoid the shortcomings of the shallow effects of roller compaction, the ric technique was proposed. two test sites were made in february 2011. a comprehensive investigation was made on site before and after compaction [19, 20]. • option 3 (es 3): soil mixing on site to achieve liquefaction resistance, low permeability and better willingness to compaction. the idea was to excavate top soil and mix it with quaternary gravel in such proportion that it would fulfil three criteria: (a) resistance against liquefaction, (b) the designed permeability to eliminate the need for additional sealing of the dike and (c) improved willingness to compaction with reduced influence of seasonal water content changes on stiffness. the trial tests were made in april and may 2014, and the roller with the roller integrated continuous compaction control (ccc) was used for compaction. as the trial fields showed promising results, the temporary, app. 1.2 km long dike was constructed in order to protect the deep excavation for the powerhouse against the floods, as well as to get more data about the homogeneity, permeability and the resistance against liquefaction of the dike. figure 8 shows the locations of the field trial tests es 1 to es 3 and the temporary dike. 4. results from the experimental sites 4.1. experimental site es 1: roller compaction trial field 4.1.1. site description the soil profile and the soil properties were the same as presented in figure 2 and table 2. the thickness of the top layer was 4 m, the roots appeared in dense network to the depth up to 3.5 m. figure 9 shows the depth of the roots (left) and the size of the first of four test fields prepared to study the roller compaction efficiency of the foundation ground – subgrade (right). figure figure 8. location of the field trial tests: es 1: roller compaction; es 2: ric; es 3 and temporary dike: soil mixing. 10 shows the test embankment (left), constructed by using the silty sands from the top layer and the amount of coal particles (right) that floated up after the layer wetting. during compaction different modes of roller operation, layer thickness, drying and wetting of compacted layers were studied. the efficiency of the roller compaction was measured by using the nuclear gauge and by the calibrated cylinders to measure the water content and the density. stiffness of compacted layers was measured by static and dynamic plate load test. undisturbed samples were taken to check the liquefaction potential in the csst. the spt, cpt and the sdmt tests, planned to be realised at the test sites, were not possible, as the flood destroyed the trial sites soon after construction. 4.1.2. results and discussion at the foundation ground, the roller compaction was found not effective due to the influence of soft sublayers, coal particles and roots. the average achieved dp r, expressed as ratio ρd/ρd,max, was 88 % and the average stiffness, measured by the dynamic load plate test (zorn) was evd < 7 mpa. the efficiency of compaction on embankment layers strongly depended on the layer position above the ground level, as the influence of the soft subgrade decreases with the embankment height. at the level of 2 m above the subgrade, the average compaction of dp r = 94 % was achieved (figure 11), while the average stiffness of the embankment layers remained at the same average ranges as measured on the subgrade, evd = 5 − 10 mpa. figure 12 compares the results of dry densities (left) and water content (right), achieved by using the nuclear gauge and the calibrated cylinders. the nuclear gauge gives significantly higher values of water content, which can be explained by the presence of coal particles and the nuclear gauge water content mea21 a. petkovšek, m. maček, j. smolar acta polytechnica ctu proceedings figure 9. depth of the roots (left) and the size of the first of four test fields prepared to study the roller compaction efficiency (right). figure 10. finished test embankment (left) and the amount of coal particles (right). suring principles. figure 7 (red diamonds) shows the csst results on the undisturbed samples from the test field. the data conform well to the preliminary forecast. 4.1.3. experimental site es 1 conclusions with the use of roller compaction, it is not possible to improve the top soil layer in order to reduce its liquefaction potential. the dynamic and static load plate test results strongly depend on saturation/water content and therefore do not measure parameters, which are decisive for the intended purpose. 4.2. experimental site es 2: rapid impact compaction (ric) test fields 4.2.1. background the hpp brežice test sites represented the first attempt to introduce the ric in slovenia. the detailed description of the ric test site is given in vukadin [20]. to highlight the importance of the test method validity and reliability, we will focus our discussion on the results of the dmt and sdmt tests, not reported in vukadin [20], as well as on the results of the cpt and dpsh b test, interpreted from a standpoint differing from that of vukadin [20]. 4.2.2. ric test site description the locations of the test sites are shown in figure 8, while a view to the ric during operation and the test site layout is given on figure 13. rough dimensions of each test field were 18 × 25 m. before the ric was put in operation, the vegetation layer of app. 1 m had been removed from the test site. the thickness of the top layer was from 3.8 to 4.0 m, with the fines content between 21 and 66 %, water content between 16 and 21 % and the dry densities of 1.2 – 1.6 t/m3 [20]. the gw table was in the quaternary gravel. at each test field, a comprehensive program of field tests (cpt, dmt, sdmt, dpsh b and dynamic cone penetrometer (dcp) panda) was performed before and after compaction. the filed measurements were complemented by samplings from trial pits and laboratory tests. each test field was divided in 4 sections (p1 to p4, figure 13) in order to study the influence of different compaction modes on the compaction efficiency. the compaction point raster density increased from section p1 to section p4 and the compaction was executed in one, two or three passes. after each compaction pass the ground was levelled by filling the holes created during compaction with the locally available sand. the decision about the end of the compaction at each point was made considering two criteria: (a) the accumulated total settlement per pas of 800 mm, or (b) the minimal depth of settlement per blow less than 10 mm. the average energy per unit cell area was 800 (p1), 1160 (p2), 1000 (p3), and 1230 kj/m2 (p4). 22 vol. 10/2017 testing methods for mechanically treated soils figure 11. comparison of standard proctor test curve (spp) and dry density achieved during trial embankment compaction. figure 12. comparison between nuclear gauge and laboratory measurements of water content and dry density. 4.2.3. results and discussion the efficiency of ric was controlled by using different field test methods. this paper will focus primarily on the interpretation of the cpt, sdmt and dpsh b results. cpt test figure 14 shows the soil behaviour type (sbtn), the normalized cone resistance (qtn) and the sleeve friction resistance (fs) evaluated from the cpt test. the qtn is a non-dimensional parameter taking into account the in-situ vertical effective stress and the soil type and is not depth dependent. the fs depends on friction between soil and metal sleeve and horizontal effective stress. the horizontal stress increases with vertical effective stress (σ′v) and could be calculated by multiplying σ′v with horizontal stress coefficient during the cpt testing. the influence of compaction on the fs is given in two diagrams, as the relationship fs – depth and fs/σ′v depth. the fs/σ′v diagram presents the increase of horizontal stress coefficient during the cpt testing due to compaction. due to the impact of levelling of ground surface, the uppermost 1 m of the test field is excluded from the evaluation of the results. figure 14 shows that classification based on sbtn was not changed due to compaction (5: sand mixtures – silty sand to sandy silt and 6: sands – clean sand to silty sand). the qtn increases from the initial 40 to 60-160 after the ric. in natural ground the fs is increasing with the depth. after compaction the fs increases slightly to the depth of 1.5 m. at the depth of 1.5 m the sharp increase of fs is detected towards the value of 100 kpa. the fs/σ′v is constant with the depth before ric, thus indicating homogeneous ground conditions. after the ric the highest increase of fs/σ′v is observed at the depth of 1.6 m and then decreases with the depth towards the initial value at the depth of 4 m. the cpt test did not recognize effects of different compaction energies on the increase of qtn and the fs. figure 15 compares the qtn and the fs before and after ric. after compaction the qtn is 2 to 4 times higher and the fs is 2 to 6 times higher than before compaction. below the depth of 3 m the effect of ric is decreased. no influence of different compaction energies used for the compaction can be recognized. (s)dmt test the results of the dmt test are presented as intermediate parameters: material index (id), horizontal 23 a. petkovšek, m. maček, j. smolar acta polytechnica ctu proceedings figure 13. ric during compaction (left) and the test site es 2 layout (right – adapted from vukadin, [20]). figure 14. results of cpt tests. cpt 1 and cpt 2 before ric, cpt p1 to cpt p4 after ric performed on sub-fields p1 to p4 (tests performed by irgo, [19]). figure 15. results of cpt after ric compared to those before ric (tests performed by irgo, [19]). 24 vol. 10/2017 testing methods for mechanically treated soils stress index (kd), and dilatometer modulus (ed) (figure 16). the id and kd are effective vertical stress independent. the value of ed is presented also in relationship with the vertical effective stress (ed/σ′v) [21]. results show that the classification based on id changed a little due to compaction (mostly sandy silt and silty sand). the general soil behaviour type remains the same. below the depth of 1 m, the kd was around 5 before compaction. this value was higher than expected for normally consolidated soils [22]. after compaction the kd first increases and then, below the depth of 2 m, decreases; however, it remains slightly higher than recorded before compaction. the kd is related to fs [23], as the first one gives a ratio between horizontal and vertical effective stress during the driving of dmt, and the second one reflects the friction between soil and metal sleeve induced by horizontal effective stress during the driving of the cpt probe. therefore, the development trend of the kd with the depth should be of the same range as the developing of fs/σ′v. the ed increases with the depth before compaction, as expected, and results in higher values after compaction. the ed/σ′v exhibits nearly constant values with the depth before compaction and clearly increases after compaction. to estimate the ric efficiency with the depth, the values of kd and ed before and after compaction are compared (figure 17). the kd increases by 2 to 3 times and the ed increases by 2 to 4 times. the dmt tests after compaction exhibit almost no increase of kd and ed at the depth of 2 to 2.5 m. at the depth of 3 m both parameters start to decrease towards the values measured before compaction; however, the decrease is not as clear as in case of cpt. contrary to the cpt results, the dmt shows clear influence of compaction energy on the increase of the measured values. the sdmt enables measurements of shear wave velocity (vs). shear wave velocity depends on void ratio, effective vertical stress, stress history, cementation and over-consolidation (ocr) ratio [24]. the increase of all parameters with the exception of void ratio increases the vs. during compaction the void ratio decreases and the ocr increases. thus, the increase of vs is expected. figure 18 compares the constrained moduli (mv /cpt and md/dmt) and the vs before and after ric. the right graph gives also the lines of vs, calculated by using the empirical equation (3) given by hardin and drnevich [25]. gmax = 1230 · (2.973 − e)2 (1 + e) · (ocr)k · p′ 1 2 (3) gmax = ρ · v2s (4) where gmax is shear modulus at small deformations, ρ is density, e is void ratio, ocr overconsolidation ratio, k is empirical material factor (depends on plasticity index) and p′ is mean effective stress. figure 18 shows that before compaction, the vs, calculated after hardin and drnevich [25], is lower than measured and after compaction the calculated vs is close to the measured. when comparing the measured values of vs before and after compaction, no improvement and no influence of the compaction energy can be detected. cpt and dmt show the increase of constrained modulus (mv and md) after compaction. the range of measured values after compaction is similar. this is not in accordance with the findings reported by schmertmann [26], lee et al. [27], among others, who found the increase of dmt constrained modulus (md) higher than that of cpt (mv ). dynamic probing super heavy (dpsh b) dpsh b raw results are presented in figure 19 (upper line). before compaction the number of blows (n20sb) increases slightly with the depth. the n20sb values are lower than minimum required (5) according to en iso 22476-2. after compaction the n20sb values are higher and nearly constant with the depth, but the scatter of the result is high. from the n20sb the equivalent spt blow-count (n1)60 was calculated using the empirical equation (5) adapted after macrobert et al. [28] and using corrections given by ec 7.2. equivalent nsp t = n20sb 0.02 · n20sb + 0.53 (5) dpsh b clearly indicates the efficiency of ric and the influence of compaction energy on the increase of the measured values. from this point of view, the sensitivity of the dpsh b test could be estimated as more indicative for the intended purpose than the sensitivity of cpt, dmt and sdmt methods. contrary to cpt and sdmt, the dpsh b confirms the compaction efficiency along the whole depth of the top layer, which was not observed by using the cpt and sdmt. the question for the discussion is: which method could be estimated as decisive? can we estimate the efficiency of the improvement by preferring optimistic test results (dpsh b) and neglecting less optimistic (cpt, sdmt) results? experimental site es 2, conclusions the analysis of the test results confirm that the ric technique is an effective tool for improving the geotechnical properties of silty and sandy soils, as it was already reported by adam and paulmichl [29]. however, the decision about the use of the ric technique to reduce the liquefaction potential must rely on the test results which are consistent and confirm without any doubts the efficiency of the improvement. as it was shown in the graphs, the scattering of the collected data raises serious questions about the validity, sensitivity and reliability of the test methods used to confirm the ric efficiency as well as questions about the homogeneity of the improvement. analysis of the results shows: 25 a. petkovšek, m. maček, j. smolar acta polytechnica ctu proceedings figure 16. intermediate dmt results. dmt 1 before ric, dmt p1 and dmt p4 after ric performed on sub-fields p1 and p4. figure 17. results of sdmt after ric compared to those before ric. figure 18. constrained modulus (mv-cpt and md-dmt)and measured/calculated vs (sdmt) 26 vol. 10/2017 testing methods for mechanically treated soils figure 19. comparison of n20sb and calculated (n1)60 values before and after compaction (tests performed by irgo, [19]). • the cpt and the dmt do not confirm the compaction efficiency along the whole depth of the top layer, while the dpsh b does • the cpt does not recognize the effect of compaction energy on the interpreted values, while the dmt and the dpsh b do • the sdmt (vs) does not recognize any improvement after compaction • the cpt and the dmt clearly recognize the less effective compaction at the depth of 2 m, while the dpsh b does not. although the dpsh b tests clearly show the efficiency of ric, the achieved improvement does not satisfy the criteria needed to reduce the liquefaction potential (figure 19). due to lack of previous experiences with the ric technique in slovenia and inconsistency of the test results, the ric technique was not recognized as a promising method for reducing the liquefaction potential. however, the authors of the paper believe that by preparing more test fields and by using slightly modified execution of ric and field testing, the results would be different and very probably more promising, too. 4.3. experimental site es 3: soil mixing 4.3.1. background the idea of soil mixing arose by joining three demands: (a) to reduce the liquefaction potential of silty-sandy soils, (b) to assure low enough permeability to avoid any additional sealing of the dikes, and (c) to improve the mixture willingness to compaction. 4.3.2. soil mixing test site description figure 8 shows the location of the test site and figure 20 gives a view to one of the test sites during test compaction and calibration of roller with ccc. as the results from the test fields were promising, an app. 1.2 km long temporary dike was constructed to protect the up to 70 m deep excavation and to gain additional data necessary to support the decision for the construction of 14 km long dikes (figure 21). 4.3.3. soil mixture demands before the soil mixing tests started, the designer prepared the requirement for the soil mixture, as follows: 27 a. petkovšek, m. maček, j. smolar acta polytechnica ctu proceedings • hydraulic conductivity lower than 1 × 10−6 m/s • shear strength must assure stable slopes at 1:1.75 (β = 30°) • homogeneity and good willingness to compaction • resistance to liquefaction, as already explained in previous chapters. 4.3.4. test program the test program was divided in five steps, i.e.: • mixture design, using the kozeny-carman equation • laboratory mixture preparation and testing • test fields on site to define the mixing and compaction procedures and to correct the laboratory mixture design, if necessary • construction of temporary dike with the qa-qc control tests during construction, using the ccc and conventional field test methods: the nuclear gauge, the dynamic and static load plate test and the double ring infiltrometer test • field testing on the temporary dike by using the sdmt, dpsh b and geohydraulic tests in boreholes. 4.3.5. the laboratory mixture design the mixture was proposed by using the kozeny–carman semi-empirical equation [30] [31] [32]. the equation relates the hydraulic conductivity to specific surface area, porosity and tortuosity. the main obstacle to wider use of the kozeny–carman equation is the specific surface area (ssa) of soil particles [33]. carrier [34], among others, proposed the estimation of ssa based on grain size distribution. the derived hydraulic conductivity based on the kozeny–carman equation and estimation of the ssa area based on grain size distribution is: k = 1.99 × 104   100 %∑ fi d0.404i−1 · d 0.595 i  ·( 1 sf )2 · ( e3 1 + e ) (6) where k is hydraulic conductivity at 20 ◦c (cm/s), sf shape factor (6 for rounded grains), di particle size at sieve size i (cm), fi fraction between sieve i and i-1, e void ratio (%). the problem of the above equation is the determination of the finest fraction. based on chapuis in legare [35], 1/3 of the smallest measured grain size could be used for the estimation of an equivalent grain size of this fraction. fi d 0.404 i−1 · d 0.595 i = fi 1 3 di−1 (7) there are some limitations to the use of the kozeny–carman equation: • not valid for clays • valid for laminar flow of water (darcy law) and therefore not valid for clean, coarse gravel with pore size greater than 3 mm and large hydraulic gradients • the grains should be rounded or cubic (blocky) • soils with a long tail of small fractions (soils which are subject to internal erosion and segregation) can make problems. particles from the tail of small fractions are not in soil matrix but rather on larger grains. in this case the grain size distribution could be corrected by using 20 % of the increase of passing by 10 times of the increase of the grain size [33]. • anisotropy of hydraulic conductivity. it is assumed that the calculated value relates to the vertical hydraulic conductivity. the grading ranges of the available soils are given in figure 3. based on the kozeny–carman equation the mixture of sandy and gravely soil was proposed in mass ratio of 20:80 to 25:75 (figure 22). 4.3.6. laboratory testing the hydraulic conductivity of sands was measured in the triaxial cell and in variable head permeameters, while gravel and mixtures were tested in a large constant head permeameter. the diameter (d) of the large constant head permeameter was 304 mm and its height (h) was approximately 600 mm (figure 23). there were 9 piezometers on 3 levels to measure the local hydraulic head (total potential). the disadvantage of large constant head permeameter is poor compaction due to thin plastic walls. the smaller constant head permeameter (d = 150 mm h = 125 mm) enabled proper compaction of soils with maximum grain size of 31.5 mm. during the gravel tests in large permeameter a non-linear decline of hydraulic head in piezometers was observed. three hydraulic conductivities were calculated (average, maximum and minimum). this problem indicates possible segregation and anisotropy of compacted gravels. the results of the tests are presented in figure 24. the calculated hydraulic conductivity using the kozeny-carman equation are in agreement with the laboratory test results. the measured value on different mixtures was up to 10 times different than calculated. the reason for this is very probably in material segregation and imperfect saturation. in general, the laboratory tests confirmed predictions of the kozeny-carman equation and the suitability of proposed mixture. 4.3.7. test fields the dimension of the test field was app. 50 m × 10 m. each test field was made of two 55 cm thick layers of mixture. at the beginning, due to the rainy weather, it was difficult to assure the homogeneity due to the high 28 vol. 10/2017 testing methods for mechanically treated soils figure 20. vibratory roller during compaction on one of the test fields (left) and a view to the control pit in one of the test layers (right). figure 21. a view to the temporary dike, constructed from soil mixture and used as part of the large scale test embankment. figure 22. calculated ranges of soil mixture to achieve the required hydraulic conductivity (k < 10−6m/s) and stiffness. water content of the silty sandy soils. the achieved dry density and stiffness of the mixture were in direct relationship with the water content and the fines content, detected at each controlling point (figure 25). on mixtures with higher content of fines the achieved layer stiffness after compaction was lower. the difference in stiffness on different test tracks was clearly observed by ccc (figure 26). later on, the weather improved and the mixing processes and the compaction ran according to expectations. the decision was accepted to construct a temporary dike by using the soil mixture. 4.3.8. temporary dike construction control during the construction of the temporary dike, a comprehensive program of control test was realised, including the 61 sieve analyses and water content tests, 13 double ring infiltrometer tests, 325 spot tests – nuclear gauge. the ccc was used to control the stiffness and homogeneity. 53 gradation curves were inside the required area and 8 samples contained app. 1% lower content of fines than required. results of all hydraulic conductivity tests were inside the design requirements (figure 27). the average achieved density was 98 % of the modified proctor (mpt) maximum dry density. 4.3.9. temporary dike post construction control two years after the construction control tests were made from the dike crown. in three bore holes field water permeability (variable head) tests were performed at different depths. two tests (at depths of 2-4 m and 4-6 m) were performed in the dike and one test (at a depth of 6-8 m) in the foundation ground (figure 27). four tests in the dike met the design requirements (1 × 10−6 m/s) and two tests indicated k higher than required. results from four dpsh b probing are presented in figure 28. as the crown has been used as a temporary unsealed road for the needs of construction site for 2 years, higher n20sb values were observed at the depths up to 2 m below the crown. at greater depths 29 a. petkovšek, m. maček, j. smolar acta polytechnica ctu proceedings figure 23. a view to the permeameter cell (left) and a view to the surface of gravel and mixture (right). figure 24. comparison of calculated (kc) and measured hydraulic conductivity (k). the n20sb values were higher than recorded at the ric test sites. three sdmt probings were performed. before the test started, the boreholes were filled with dry uniform sand. therefore, a greater variation in result can be expected due to the effects of backfill. the third probe (sdmt3b) was repeated to check the repeatability of the results due to the difficulties during the test in probe sdmt3. the vs results in all probes at the dike were around 200 to 450 m/s and were significantly higher than those measured in the top layer before compaction and after the ric compaction. the measured shear wave velocities are close to the values calculated by using hardin and drnevich equation [25] (figure 28). 4.3.10. experimental site es 3 conclusions the kozeny-carman equation was found to be an effective tool for the design of the mixture with the given hydraulic conductivity considering the equation limitations. the laboratory test and the field double ring infiltrometer test results were in good agreement with the calculation. however, two of six tests in the borehole using open systems gave by a decade higher value of k. the ccc test results show that the compaction behaviour and layer stiffness of sandy gravel mixtures containing app. 10 20 % of fines are mainly controlled by the water content of fines. this means that even though the mixture is classified as silty gravel, its compaction behaviour is typical for less permeable soil mixtures. the achieved dry density after compaction, the sdmt (vs) and the dpsh b tests show that the mixture has properties of gravel. figure 29 shows that for a given csr the mixing of locally available soil is a promising method to reduce the liquefaction potential. however, the decision was made not to use the soil mixing for the construction of the dikes. all the soils from the top layer were removed and replaced by clean gravel. gravel was also used as fill material for all dikes and the sealing of the dikes was achieved by the geosynthetic bentonite barriers. 30 vol. 10/2017 testing methods for mechanically treated soils figure 25. results from the first test section. high water content in silty part of the mixture control the achieved stiffness and density. evd – dynamic deformation modulus measured by dynamic load plate test; evib – dynamic elasticity modulus of soil beneath drum (terrameter ccc system). figure 26. ccc record from one of the first test sections showing heterogeneity of the compaction. figure 27. comparison of hydraulic conductivity predicted after kozeny-carman equation (upper limit, lower limit) and measured values in laboratory and on site. 5. conclusion mechanical stabilisation for ground improvement has become a widely accepted system to support foundations for a wide variety of structures as well as for the construction of embankments and embankment dams. the use of mechanically stabilized locally available soils is also one of the top priorities of the philosophy generally understood as “sustainable development” [36]. the paper focuses on the interpretation of the results of the tests used for the control of the efficiency of some of the stabilisation techniques. the analysis of the tests shows inconclusive results. in foundation ground, where the requirements are different than those given for regular earthworks and the safety demands are higher, the use of different test methods is necessary. due to the lack of experiences, the use of different methods may lead to inconsistent results and often raises more questions than gives answers. understanding the local conditions and mechanisms that may influence the test results in different soil types/stabilisation techniques is important for engineering judgement and leads to the ability to make tough decision. however, past experiences are also important and sometimes a lot of time and energy are needed before a new method/ technique becomes widely accepted. 6. acknowledgement this work was funded by the hess and realised in close cooperation with irgo ljubljana. authors would like to express special thanks to mr. b.sc. andrej unetič (hse invest) for his help at realising the investigation program and to many engineers and technicians for their essential field work and data collection. the advice of prof. dietmar adam from tu vienna is highly appreciated. 31 a. petkovšek, m. maček, j. smolar acta polytechnica ctu proceedings figure 28. temporary dike. shear wave velocity, hydraulic conductivity, and results of dpsh b (from left to right). figure 29. diagrams for evaluation of liquefaction potential based on (n1)60 (upper) and vs1 (lower). bc – before compaction, fc – fines content, mix – soil mixture, ric – rapid impact compaction, ts – top soil layer. references [1] w. f. v. impe. soil improvement techniques and their evolution. aa balkema, rotterdam, netherlands., 1989. [2] h. brandl. geotechnics for roads, rail tracks and earth structures, chap. the importance of optimum compaction of soil and other granular material, pp. 47–68. crc press, 2001. [3] i. vaníček, m. vaníček. earth structures in transport, water and environmental engineering. springer netherlands, 2008. [4] v. r. raju, j. daramalinggam. ground improvement: principles and applications in asia 165(2). [5] t. verbič. quaternary stratigraphy and neotectonics of the eastern krško basin part 2: neotectonics. razprave iv. razreda sazu, xlvi-1, 2005. [6] v. ribarič. seizmičnost slovenije – katalog potresov. seizmološki zavod sr slovenije, ljubljana, 1982. [7] ž. veinović, d. domitrovič, t. lovrić. historical occurrence of liquefaction in zagreb area and estimation of reoccurrence in case of another strong earthquake. rudarsko-geološko-naftni zbornik 19:111–120, 2007. [8] ž. veinović, p. kvasnička, v. jurak. geohazard characteristics of the sava river plain near zagreb. konferencija razvitak zagreba 2008. acessed 15.9.2016. [9] i. idriss, r. boulanger. semi-empirical procedures for evaluating liquefaction potential during earthquakes. soil dynamics and earthquake engineering 26(2):115– 130, 2006. 11th international conference on soil dynamics and earthquake engineering (icsdee): part ii, doi:https://doi.org/10.1016/j.soildyn.2004.11.023. [10] d. aleksovski, g. mirakovski, d. dojčinovski, et al. he brežice geofizikalne raziskave in seizmična mikrorajonizacija,. geofizikalne raziskave in seizmična mikrorajonizacija, iziis univerze sv. kiril i metodij, skopje., 2008. [11] arso, slovenian environment agency. agencija republike slovenie za okolje. 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[16] t. l. youd, s. k. noble. liquefaction criteria based on statistical and probabilistic analyses p. 201–215, 1997. [17] r. d. andrus, k. h. stokoe. liquefaction resistance of soils from shear-wave velocity. journal of geotechnical and geoenvironmental engineering 126(11):1015–1025, 2000. doi:10.1061/(asce)1090-0241(2000)126:11(1015). [18] s. l. kramer. geotechnical earthquake engineering. prentice-hall international series in civil engineering and engineering mechanics. upper saddle river, new jersey., 1996. [19] irgo. he brežice, poročilo o impulznem globinskem zgoščevanju na testnih poljih. faza: pgd. report nr. ic161/11. ljubljana, slovenia., 2011. [20] v. vukadin. the improvement of the loosely deposited sands and silts with the rapid impact compaction technique on brežice test sites. engineering geology 160(supplement c):69–80, 2013. doi:https://doi.org/10.1016/j.enggeo.2013.03.025. [21] p. k. robertson. cpt-dmt correlations. journal of geotechnical and geoenvironmental engineering 135(11):1762–1771, 2009. doi:10.1061/(asce)gt.1943-5606.0000119. [22] s. marchetti. in situ tests by flat dilatometer. journal of the geotechnical engineering division 106(3):299–321, 1980. [23] j. m. o. hughes, p. k. robertson. full-displacement pressuremeter testing in sand. canadian geotechnical journal 22(3):298–307, 1985. doi:10.1139/t85-043. [24] r. dobry, m. vucetic. state of the art report: dynamic properties and seismic response of soft clay deposits. in proceedings of the intl. symposium on geotechnical eng. of soft soils, vol. 2, pp. 51–87. mexico city, 1987. [25] b. hardin, v. p. drnevich. shear modulus and damping in soils: design equations and curves 98:667–692, 1972. [26] j. h. schmertmann. comparing dmt with cpt in nc/oc sand bucket test. dmt digest (4), 1984. [27] m. lee, s. choi, m. kim, w. lee. effect of stress history on cpt and dmt results in sand. engineering geology 117(3-4):259–265, 2011. doi:10.1016/j.enggeo.2010.11.005. [28] c. macrobert, d. kalumba, p. beales. correlating standard penetration test and dynamic probe super heavy penetration resistance values in sandy soils. journal of the south african institution of civil engineering 53(1):46–54, 2011. [29] d. adam, i. paulmichl. rapid impact compactor an innovative dynamic compacttion device for soil improvement. in proc. 8th international geotechnical conference, pp. 183–192. slovak university of technology, bratislava, 2007. [30] j. kozeny. über kapillare leitung des wassers im boden, 136 (2a). hölder-pichler-tempsky, a.-g. akad. d. wiss., wien., 1927. [31] p. c. carman. the determination of the specific surface of powders. journal of the society of chemical industry 57:225–238, 1938. [32] p. c. carman. flow of gases through porous media. new york: academic press, 1956. [33] r. chapuis, m. aubertin. predicting the coefficient of permeability of soils using the kozeny-carman equation. école polytechnique de montréal 2003. [34] w. d. carrier. goodbye, hazen; hello, kozeny-carman. journal of geotechnical and geoenvironmental engineering 129(11):1054–1056, 2003. doi:10.1061/(asce)1090-0241(2003)129:11(1054). [35] r. chapuis, p. p. légaré. a simple method for determining the surface area of fine aggregates and fillers in bituminous mixtures. effects of aggregates and mineral fillers on asphalt mixture performance pp. 177–177–10, 1992. doi:10.1520/stp24217s. [36] i. vaníček. sustainable construction. ed. 1st. czech technical university in prague, prague, 2011. 33 http://dx.doi.org/10.3208/sandf1972.20.4_127 http://dx.doi.org/10.1061/(asce)1090-0241(2000)126:11(1015) http://dx.doi.org/https://doi.org/10.1016/j.enggeo.2013.03.025 http://dx.doi.org/10.1061/(asce)gt.1943-5606.0000119 http://dx.doi.org/10.1139/t85-043 http://dx.doi.org/10.1016/j.enggeo.2010.11.005 http://dx.doi.org/10.1061/(asce)1090-0241(2003)129:11(1054) http://dx.doi.org/10.1520/stp24217s acta polytechnica ctu proceedings 10:16–33, 2017 1 introduction 2 background 2.1 site description 2.2 geological profile 2.3 soil properties 2.3.1 index properties 2.3.2 liquefaction potential 2.4 preliminary conclusions 3 discussion on improvement technique 4 results from the experimental sites 4.1 experimental site es 1: roller compaction trial field 4.1.1 site description 4.1.2 results and discussion 4.1.3 experimental site es 1 conclusions 4.2 experimental site es 2: rapid impact compaction (ric) test fields 4.2.1 background 4.2.2 ric test site description 4.2.3 results and discussion 4.3 experimental site es 3: soil mixing 4.3.1 background 4.3.2 soil mixing test site description 4.3.3 soil mixture demands 4.3.4 test program 4.3.5 the laboratory mixture design 4.3.6 laboratory testing 4.3.7 test fields 4.3.8 temporary dike construction control 4.3.9 temporary dike post construction control 4.3.10 experimental site es 3 conclusions 5 conclusion 6 acknowledgement references acta polytechnica ctu proceedings doi:10.14311/app.2017.9.0013 acta polytechnica ctu proceedings 9:13–18, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app influence of machining technologies on values of residual stresses of oxide cutting ceramics jakub němeček∗, kamil kolařík, jiří čapek, nikolaj ganev department of solid state engineering, faculty of nuclear science and physical engineering, czech technical university in prague ∗ corresponding author: j.nemecek91@gmail.com abstract. currently, the intensive development of engineering ceramic and effort to replace sintered carbides as cutting materials are in progress. with the development of the sintering technology it is now possible to produce compact ceramic cutting samples with very good mechanical properties. the advantage of these materials is their easy accessibility and low purchase price. in this work, the influence of the finishing machine technology on the values of surface residual stresses of cutting ceramic samples al2o3+tic were studying. the samples were supplied by moscow state university of technology stankin. measurements made in the x-ray diffraction laboratory at the department of solid state engineering were performed for both the phases. the influence of the parameters of machining to residual stresses was studied and the resulting values were compared with each other. keywords: cutting ceramic; machining technology; x-ray diffraction; residual stresses. 1. introduction ceramic tools have been used since the stone age, when sandstone containing large amounts of silica blades was used, for example for the sharpening of knives, razor blades or other tools. this material was used henceforward as abrasive wheels during the ages of cutting weapons for sharpening of swords, knives and other blades. only recently these grinding tools have been replaced by other modern materials for example by sic or diamond. at the beginning of the 20th century, the first use of al2o3 ceramics as a cutting tool occurred. over the years, its development had moved to a level at which we were able to replace, for example, sintered carbides using those ceramic materials. this effort has been caused by the mechanical properties of aluminium oxide and the easy availability of the starting materials [1, 2]. the aim of this paper is to study the real structure of al2o3+tic oxide cutting ceramic. the microstrains and a size of crystallites was measured and influence of machining technologies on the macroscopic residual stresses was studied. 1.1. preparation of samples the starting materials and their properties have been known for a long time. the problem was how to make compact samples with the desired mechanical properties and dimensions. the current production process consists of three parts: first, grinding the starting material to a powder with grain size of the order of 100 nm, second, sintering the powders into a compact form and third, the surface treatment of the samples to the desired dimensions. an important technological aspect of the properties of cutting ceramics is the grain size of the starting powders. if the individual grain size is approximately the same (if they have ideal spherical shape), the uniform distribution and the pore size are gained more easily during forming and thereby for uniform compaction throughout the volume and easier reorganization of the particles during sintering. preference is also to achieve the smallest grain size of the sample, because it leads to faster material transport [3]. the sintering process is very similar to that used for sintered carbides. the powder is pressed below the melting point into a compact body, see figure 1. but the main difference is the absence of a binder which during sintering formed a liquid phase, because it would cause degradation of mechanical properties. today, ceramic is most commonly sintered into rods of cross section of finite samples and then cut to the desired dimensions. individual plates are then variously surface treated and shaped [4]. 2. x-ray diffraction analysis of macroscopic residual stresses in cooperation with the moscow state technological university stankin that provides manufacturing technology, the influence of machining technologies on the value of residual stresses in samples of al2o3+tic oxide cutting ceramics were investigated in x-ray diffraction laboratory of the department of solid state engineering of fnspe ctu in prague. all experimental errors were around 10 % of absolute value of measurement. 2.1. grinding by diamond disc after sintering, all samples were ground using a diamond wheel. an effect of displacement of the workpiece on the lift at constant cutting speed, feed rate 13 http://dx.doi.org/10.14311/app.2017.9.0013 http://ojs.cvut.cz/ojs/index.php/app j. nemecek, k. kolarik, j. capek, n. ganev acta polytechnica ctu proceedings figure 1. scheme of creation technology of cutting ceramic sample: refinement powders (wet grinding), evaporation, pressing, sintering [5]. sample fd, displacement on the lift [mm/displ.] ap, depth of cut [mm] 1 0.5 0.04 2 1.0 0.04 3 1.5 0.04 table 1. parameters of grinding by diamond disc. figure 2. scheme of grinding by diamond disc [5]. of the workpiece in the cutting direction and depth of cut was observed. grinding parameters are given in table 1, where cutting speed vc = 30 m/s and feed rate f = 12 m/min. values of residual stresses were determined in two orthogonal directions l (direction of the grinding wheel) and t (direction of displacement on the lift). because of unidirectional grinding anisotropy of values was observed in both directions. different values are due to the different structure of al2o3 (trigonal) and tic (cubic) and the different number of slip planes (see table 2). 2.2. thermal annealing of the samples after grinding by diamond disc next samples were ground by diamond disc with the same machining parameters and then thermally anfigure 3. structure of al2o3 (up) and tic (down). nealed at 800 ◦c. it is evident from table 3 that the values of the macroscopic residual stresses decreased considerably. in contrast, the value of microstrains and grain size given by the parameter fwhm remained almost unchanged. 2.3. air abrasive machining in this method of machining, air of high pressure with added cutting microparticles is jetted on a rotating sample. the effect of air pressure changes on the residual stress in the sample was studied (table 4) with constant time of blasting t = 60 s and rotates per minute rpm = 50. the obtained values (table 5) show that when a pressure increases from 2 to 2.5 bars a decrease of compressive residual stresses (stress relaxation) is ob14 vol. 9/2017 influence of machining technologies on values of residual stresses al2o3 tic l direction rs mnkα (0210) l direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 1 -216 33 0.787 1 -518 77 0.759 2 -182 39 0.693 2 -536 77 0.741 3 -165 37 0.859 3 -530 74 0.708 t direction rs mnkα (0210) t direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 1 -436 36 0.804 1 -776 67 0.668 2 -470 30 0.709 2 -805 93 0.759 3 -580 37 0.772 3 -938 89 0.679 table 2. the values of normal (σn ) and shear (σs) residual stresses after grinding by a diamond disc. al2o3 tic l direction rs mnkα (0210) l direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 12 -185 19 0.666 12 -516 91 0.965 13 -203 16 0.888 13 -290 40 0.609 14 -199 30 0.889 14 -289 112 0.847 t direction rs mnkα (0210) t direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 12 -303 20 0.781 12 -599 71 0.782 13 -359 6 0.802 13 -474 57 0.613 14 -342 27 0.769 14 -473 104 0.908 table 3. the values of normal (σn ) and shear (σs) residual stresses after thermal annealing. sample p, air pressure [mpa] 4 1.5 5 2.0 6 2.5 table 4. parameters of air abrasive machining. figure 4. scheme of air abrasive machining and a picture of machining chamber [5]. served. it is probably a result of overcoming the yield strength of the material. perhaps machining process with higher pressure values influences deeper sub-surface layers of material. the energy required for crack propagation is therefore higher for these seemingly lower pressure residual stresses. to confirm this assertion, a depth-profile analysis of residual stresses by sequential force-free etching of surface layers is needed. rotational movement of the specimen during the machining leads to an isotropic stress distribution i.e., the values in both the directions l and t are similar. 2.4. water jet machining water jet machining (wjm) is based on the same principle as air abrasive machining, when a focused beams of water is jetted out by varying the pressure on the rotating sample (see figure 6 and table 6). other parameters: time of blasting t = 60 s, distance of samples l = 2 mm, number of nozzles n = 50. it is evident, that for applied values of pressure jetting the yield strength value have not been overcome. the values of residual compressive residual stresses increase steadily with increasing pressure in contrast to the samples after air abrasive machining. due to the rotating of the samples an isotropic residual stress distribution is achieved again (see table 7). 15 j. nemecek, k. kolarik, j. capek, n. ganev acta polytechnica ctu proceedings al2o3 tic l direction rs mnkα (0210) l direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 4 -352 34 0.787 4 -799 117 0.965 5 -367 45 0.694 5 -928 111 0.609 6 -242 32 0.827 6 -278 83 0.847 t direction rs mnkα (0210) t direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 4 -295 42 0.781 4 -737 119 0.641 5 -389 44 0.802 5 -899 123 0.890 6 -284 36 0.769 6 -283 69 0.611 table 5. the values of normal (σn ) and shear (σs) residual stresses after air abrasive machining. sample p, pressure [mpa] 7 1.5 8 2.0 9 2.5 table 6. parameters of water jet machining. figure 5. scheme of water jet machining. 2.5. laser shock peening during machining by laser the samples are heated and it results in stress relaxation. increased power (see table 9) will lead to higher temperatures, and thus to higher stress relaxation. furthermore, it is possible to observe the influence of the laying of individual spots (overlapping) on the anisotropy of distribution of residual stresses. 2.6. stress tests — the effect of friction after grinding of samples by diamond disc their surface was loaded by rotating steel disk at different time intervals. the influence of time of friction on values of residual stresses was investigated (see table 10). figure 6. scheme of linear and surface laying of spots for laser shock peening [5]. 3. microstrains and crystallite size simultaneously with the measurement of the macroscopic residual stresses, determination of crystallite size and microstrains was provided using the rietveld analysis. all results are demonstrated in [6]. from the measured values is evident that crystallite size after thermal annealing and after laser machining increased, which is probably due to recrystallization during these treatments. a decrease of microstrains is observed after air abrasive machining when the yield point was overcome and the grain sizes increased. after stress tests by steel disc the largest crystallite size al2o3 are observed, that could be again due to thermal relaxation and recrystallization of the surface. 4. conclusions values of residual stresses of cutting ceramics al2o3+tic were analysed in two perpendicular directions. the influence of machining technologies to residual stresses was studied in both the phases. anisotropy and different values for the two phases during grinding by diamond disc were obtained, and 16 vol. 9/2017 influence of machining technologies on values of residual stresses al2o3 tic l direction rs mnkα (0210) l direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 7 -99 38 0.712 7 -406 118 0.644 8 -149 42 0.733 8 -363 74 0.691 9 -255 47 0.732 9 -568 118 0.765 t direction rs mnkα (0210) t direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 7 -85 43 0.790 7 -374 116 0.644 8 -166 50 0.752 8 -364 88 0.689 9 -238 45 0.668 9 -484 108 0.669 table 7. the values of normal (σn ) and shear (σs) residual stresses after water jet machining. al2o3 tic l direction rs mnkα (0210) l direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 10.1 -53 31 0.731 10.1 -318 111 0.594 10.2 31 31 0.717 10.2 -259 96 0.657 10.3 -29 23 0.757 10.3 -200 106 0.637 10.4 -62 17 0.751 10.4 -220 95 0.568 t direction rs mnkα (0210) t direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 10.1 -155 9 0.856 10.1 -509 148 0.619 10.2 -125 9 0.868 10.2 -448 135 0.545 10.3 -71 42 0.737 10.3 -367 135 0.543 10.4 -48 39 0.704 10.4 -288 123 0.618 table 8. the values of normal (σn ) and shear (σs) residual stresses after laser shock peening. sample p, power [w] 10.1 2 10.2 5 10.3 10 10.4 15 table 9. parameters of laser shock peening. pulse frequency f = 30 khz, displacement sx = sy = 40 µm. the thermal relaxation of macroscopic residual stresses after thermal annealing was also measured. for air abrasive cutting overcoming of the yield stress was found out with increasing the pressure to 2.5 mpa. due to rotational movement during air abrasive machining and water jet machining the values of surface residual stresses are isotropic. samples machined by laser exhibit thermal relaxsample t, time of loading [s] 11.1 2 11.2 6 11.3 12 table 10. parameters of stress tests. ation of residual stresses. at the same time, the found anisotropy of state of residual stresses is due to overlapping of laser-processed areas. on the samples after the loading tests thermal relaxation of residual stresses was observed. it was shown that titancarbide is more sensitive to load than aluminium oxide. for more detailed information see [6]. acknowledgements this work was supported by grant student grant competition ctu no. sgs16/246/ohk4/3t/14 and grant agency of the czech republic (no. 14-36566g). 17 j. nemecek, k. kolarik, j. capek, n. ganev acta polytechnica ctu proceedings al2o3 tic l direction rs mnkα (0210) l direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 11.1 -42 27 0.693 11.1 -453 64 0.655 11.2 3 21 0.606 11.2 -129 76 0.530 11.3 -78 40 0.596 11.3 -242 49 0.577 t direction rs mnkα (0210) t direction rs crkα (311) sample σn[mpa] σs [mpa] fwhm [2θ] sample σn [mpa] σs [mpa] fwhm [2θ] 11.1 -145 17 0.612 11.1 -348 95 0.580 11.2 -15 19 0.610 11.2 -190 59 0.563 11.3 47 24 0.649 11.3 -240 66 0.446 table 11. the values of normal (σn ) and shear (σs) residual stresses after stress testing. references [1] r. kumar, et al. synthesis and characterization of al2o3+tic nano-composite by spark plasma sintering. international journal of refractory metals and hard materials 577:9–15, 2016. [2] z. yin, et al. friction and wear behaviors of al2o3+tic micro-nano-composite ceramic sliding against metals and hard materials. international journal of refractory metals and hard materials 577:9–15, 2016. [3] a. humár. materiály pro řezné nástroje. first printing. mm publishing, 2008. [4] c. b. carter, m. g. norton. ceramic materials: science and engineering. second edition. springer science and business media, 2013. [5] m. university stankin. materials for experimental measurement. [6] j. němeček. studium reálné struktury oxidových řezných keramik. 18 acta polytechnica ctu proceedings 9:13–18, 2017 1 introduction 1.1 preparation of samples 2 x-ray diffraction analysis of macroscopic residual stresses 2.1 grinding by diamond disc 2.2 thermal annealing of the samples after grinding by diamond disc 2.3 air abrasive machining 2.4 water jet machining 2.5 laser shock peening 2.6 stress tests — the effect of friction 3 microstrains and crystallite size 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.18.0066 acta polytechnica ctu proceedings 18:66–71, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app experimental investigation of the failure behaviour of polypropylene compounds for instrumented puncture tests florian kiehasa, ∗, anna kalteisa, michael jerabekb, zoltán majora a institute for polymer product engineering, altenberger straße 69, 4040 linz, austria b borealis polyolefine gmbh, sankt-peter-straße 25, 4020 linz, austria ∗ corresponding author: florian.kiehas@jku.at abstract. instrumented puncture tests according to iso 6603-2 and astm d3763 were executed for five different polypropylene compounds (talcum-, glass fibreand elastomer modified) with specimen thicknesses ranging from 1 mm to 4 mm. over 1500 puncture tests were executed at the impact & long-term behaviour laboratory of the company borealisr in linz. this serves as strong foundation for statistical evaluations of the ductile/brittle transition temperature. for different materials and ductile/brittle transition determination methods, similar trends have been observed, which were characterized by introducing shift factors. keywords: polypropylene, compound, puncture, mechanical testing, ductile/brittle transition. 1. introduction for polymers, the energy absorption capacity strongly depends on the morphology of the plastic and the experimental environment. in the case of impact or puncture tests, temperature and strain rate variations have a significant influence on the failure behaviour: almost all unoriented plastics exhibit brittle behaviour at low temperatures as well as at high strain rates [1]. mechanical measurements of different puncture test standards result in distinct force-deflection curves, which in turn give different transitions from ductile to brittle. furthermore, there is no concise method for the calculation of an exact transition temperature. it is assumed to be somewhere in the transition regime, a temperature range where sudden drops in energy absorption capacity indicate a change of impact behaviour from ductile to brittle. first, a precise method for the calculation has to be introduced. then, the difference of the transition as a function of specimen thickness, temperature, test standard and other parameters can be determined. 2. experimental investigation 2.1. test parameters the test setup and parameters of iso 6603-2 and astm d3763 vary greatly. the main difference is the strikerand support ring geometry. for this reference analysis, tests for both standards were executed clamped, lubricated and at a speed of 4400 mm/s to allow for comparison. therefore, the astm d3763 was slightly modified. a summary of all parameters can be looked up in table 1 and figure 1 shows the different geometries of each test standard. a more detailed description of each test standard can be looked up in literature [2, 3]. iso modified* astm specimen dimensions 60×60 mm ∅102 mm specimen thickness 1,2,3 mm 1,2,3,4 mm* striker diameter 20.0 mm 12.7 mm support ring diameter 40.0 mm 76.0 mm test speed 4400 mm/s 4400 mm/s* clamped yes yes lubricated yes yes* table 1. selected test conditions and parameters. figure 1. striker and support ring geometries iso (left) and astm (right). 2.2. materials and test schedule an extensive test program with a total number of 1586 specimens was carried out. the workload was divided between two machines: a servohydraulic roell/amsler high speed and a gravitative (spring driven) ceast fractovis plus. the comparability of measurement data obtained by both machines is documented by [4]. five different polypropylene grades developed and produced by the company borealisr were selected for this analysis. three materials are in-reactor made heterophasic copolymers, which are different in rubberand matrix design (elastomer modified polypropylene m1 to m3). m5 is also an in-reactor made heterophasic copolymer filled with talc. m4 is a glass filled polypropylene. 66 http://dx.doi.org/10.14311/app.2018.18.0066 http://ojs.cvut.cz/ojs/index.php/app vol. 18/2018 investigation of pp-compounds for instrumented puncture tests four specimen thicknesses from 1 mm to 4 mm were tested with the exception of 4 mm iso 6603-2 specimens, for which no injection moulding tool was available for manufacturing. 1 mm thick glass-fibre reinforced iso plaques were also not available. examined conditioning temperatures range from -70 ◦c to 23 ◦c. eight temperature levels with five repetitions respectively were executed for each material, test standard and specimen thickness. in certain cases, additional measurements with three repetitions were carried out in the transition regime. altogether, there are 34 test series with descriptions following table 2: polypropylene compound test standard thickness m1(elastomer modified pp) iso(iso 6603-2) t1(1mm) m2(elastomer modified pp) astm(astmd3763) t2(2mm) m3(elastomer modified pp) t3(3mm) m4(glass fibre reinforced pp) t4(4mm) m5(talcum reinforced pp) table 2. nomenclature of test series. for example, m4astmt3 represents a test series of material m4 with 3 mm thick specimens tested under the modified astm d3763 standard conditions. 2.3. data evaluation for both test standards, the failure criterion is 50 % force drop after peak load fm. the corresponding puncture energy ep is obtained by integrating the force up to the deflection at puncture sp. the transition from elastic to elastic-plastic material behaviour was described with the introduction of a yield equivalent point (sy/fy)[5, 6]. the initial slope klin (corresponding to the relative stiffness [3]) of the curve can be approximated through the specification of another characteristic data point (sy0/fy0), which is necessary to avoid deviations caused by discontinuities at the start of impact. these points are defined as having the largest orthogonal distance from the direct connection line between origin (0/0) and peak load (sm/fm). thus, the energy em expended up to the maximum load can be split into an elastic part eel and a plastic part epl. the graphical determination of the starting and end point of the initial slope, as well as the geometric interpretation of the plastic and elastic energy contributions are illustrated in figure 2. following characteristic values are examined: • maximum force/deflection/energy fm/sm/em • puncture force/deflection/energy fp/sp/ep • initial force/deflection fy0/sy0 • yield equivalent force/deflection fy/sy • elastic/inelastic energy portions eel/epl • relative stiffness klin • width of plastic plateau ∆sw klin = fy − fy0 sy − sy0 (1) eel = f 2m 2klin (2) epl = em − eel (3) figure 2. evaluation of force-deflection data. two characteristic energy ratios have been introduced to describe the elastic properties as well as the failure behaviour of force-deflection data: the elastic energy ratio rel and the ductility index [7] rduc are specified in the following equations: rel = epl + eel eel (4) rduc = ep − em ep (5) 2.3.1. curve characterization the iso 6603-2 curve type evaluation is a common method to characterize the force/deflection data obtained by puncture test ecperiments. it provides four typical types of curve progression that can usually be observed during data evaluation. in general, these range from ductile to brittle and focus on information like yielding, crack initiation and crack propagation [2]. obtained data has been divided into these four categories, which are compiled in figure 3. the determination of the ductile/brittle transition by means of curve type characterization is described in section 2.4.5. 2.3.2. optical characterization the different types of fracture appearance are taken from a company-internal standard developed by general motors corporation. it grades fracture appearances from 1 to 11 and covers almost all kinds of break that may occur during impact testing. a subdivision ranging from ductile, to semi-ductile, up to brittle can be made. illustrations of the individual fracture types are depicted in figure 4. the optical characterization was utilized to determine the ductile/brittle transition as well, which is described in section 2.4.6. 67 f. kiehas, a. kalteis, m. jerabek, z. major acta polytechnica ctu proceedings figure 3. curve types: yd yielding with deep drawing (top left), ys yielding with stable crack (top right), yu yielding with unstable crack (bottom left), ny no yielding (bottom right) [2]. 1 2 3 4 5 6 7 8 9 10 11 figure 4. fracture appearances: from ductile (1 5), to ductile-fragile (6 7), to fragile-ductile (8 9), up to fragile (10 11). 2.4. ductile/brittle transition temperature various methods for the determination of the ductile/brittle transition temperature have been proposed in literature. for impact tests with steep inclines in the transition regime, not only the tangent formation of the upperand lower shelf impact strengths [8, 9] but also optical investigations of the fracture appearance was investigated [10]. in [11] the inflection point of the temperature-dependent course was determined. likewise, the penetration energy was proposed for exploration. for this reference analysis, six different methods for the determination of the ductile/brittle transition temperature have been introduced and will be listed in the following subsections. it should be pointed out, that values are interpolated between data points. figure 5 depicts graphic representations for the determination of each transition temperature. 2.4.1. puncture energy the arithmetic mean of the local minimum and maximum puncture energy ep is calculated and the transition is assumed at the corresponding condition temfigure 5. determination of transition temperatures. perature t e. 2.4.2. elastic energy ratio the peak load energy em is compared to the elastic energy eel. it should be pointed out that the minimum value for this energy ratio is 1 when the plastic portion disappears. this case corresponds to absolutely no yielding before peak load with the corresponding curve type ny no yielding. threshold values of 1.4, 1.2 and 1.0 were examined, with t rat1.2 giving the most consistent results. 2.4.3. ductility index in contrast to the elastic energy ratio rel, the ductility index rduc is only obtained from information of the force deflection data after peak load. therefore, yielding before peak load can not be detected with this method and is neglected. a ductility index of 0 implies, that the peak load is directly followed by an unstable crack, and consequently brittle fracture. threshold values of 40 %, 30 % and 5 % were examined, with t duc30 giving the most consistent results. 2.4.4. deflection for completely brittle fractures, all characteristic events of a puncture test happen in a narrow time frame. the deflection difference between sm, sp and stotal (as well as sy when there is no plastic plateau) becomes very small. t s is the temperature where sm and sp converge. 2.4.5. curve characterization this method for the ductile/brittle transition temperature determination takes advantage of the elaborate iso6603-2 curve characterization introduced in section 2.3.1. the curve types usually occur in the sequence of yd, ys, yu and ny from high to low 68 vol. 18/2018 investigation of pp-compounds for instrumented puncture tests temperatures. an attempt was made to describe the ductile/brittle behaviour by looking out for the transitions from yu yielding with unstable crack (bottom left) to ny no yielding. this leads to two specific temperatures ty u and tn y , which are triggered at the first appearance of their respective curve type. 2.4.6. optical characterization tgm corresponds to the test temperature, where a fragile fracture type occurs for the first time. 2.5. overview transition temperatures a compilation of all calculated transition temperatures is given in table 3. with the exception of the glass fibre reinforced material m4, all materials follow similar trends between test standards and specimen thicknesses. reproducibility, consistency and comparability are the main factors that need to be considered when evaluating the transition temperatures. at the same time, the calculations involved should be as simple as possible. t duc30 shows the most inconsistent results because it is sensitive to noise of the force-deflection data after crack initiation. despite occasional outliers, the iso characterization has proven itself to be an adequate representation of the transition regime with most of the other transition temperatures lying between the boundary values of ty u and tn y . for thin specimens, tn y and ty u stretch over a wide temperature range, which gets more narrow for thicker specimens. in comparison to all the other categories, the fracture appearance usually gives much higher transition temperatures tgm , because fractures are classified as brittle as soon as crack initiation is observed disregarding any plasticity that might be present before. 2.6. transition shift factors since the trends of all categories are strikingly similar as can be seen in table 3, the shift factors are contemplated to be universally applicable for all transition temperatures. however, the impact behaviour of glass fibreand talcum reinforced pp-compounds is vastly different from the other tested materials. consequently, only elastomer modified pp-grades (m1, m2, m3) as well as the most consistent transition temperatures (t e, t rat1.2 , t s) are used to determine the shift factors. for data extrapolation, 1 mm iso test series are chosen as initial reference values, since 1 mm test series show the least standard deviations and iso is generally more consistent than astm. this procedure is illustrated in table 4. finally, these shift factors are summarized in average values α, β, γ, δ, � and ζ, which are listed in table 5. it should be pointed out, that the transition from ductile to brittle was not fully covered for the material m2, even at -70 ◦c, which is the reason why t s values for m2isot2 and m2isot3 are assumed to be -70 ◦c. t e t rat1.2 t s t duc30 tn y ty u tgm m1isot1 -47 -50 -49 -35 -52 -30 -30 m1isot1 -57 -59 -57 -53 -58 -52 -50 m1isot3 -56 -56 -58 -55 -60 -56 -56 m1astmt1 -40 -29 -30 18 -40 -23 -30 m1astmt2 -47 -45 -50 -41 -50 -40 50 m1astmt3 -43 -43 -50 -36 -50 -50 -40 m1astmt4 -45 -45 -50 -41 -50 -50 -50 m2isot1 -50 -55 -50 -40 0 -20 m2isot2 -63 -66 -70 -56 -30 -30 m2isot3 -64 -65 -70 -40 -30 m2astmt1 -40 -36 -37 7 23 -20 m2astmt2 -55 -55 -60 -41 -40 -20 m2astmt3 -52 -50 -60 -43 -50 m2astmt4 -54 -60 -22 -30 -30 m3isot1 -19 -20 -17 6 -30 -10 -20 m3isot2 -28 -29 -30 -8 -40 -10 -30 m3isot3 -26 -27 -25 -17 -34 -20 -30 m3astmt1 -23 -3 -10 16 -20 23 -10 m3astmt2 -26 -14 -19 5 -20 0 -10 m3astmt3 -26 -13 -19 7 -20 -10 -10 m3astmt4 -24 -28 -6 -30 -10 -20 m4isot2 -18 m4isot3 -7 -7 m4astmt1 -17 -17 -11 m4astmt2 -7 -27 -52 -27 m4astmt3 -6 -14 -5 -1 m4astmt4 5 -59 -22 6 m5isot1 -35 -35 -27 -4 -36 -10 -30 m5isot2 -36 -36 -36 -23 -38 -30 -40 m5isot3 -39 -36 -21 -50 -20 -40 m5astmt1 -35 -25 -21 11 -30 23 -30 m5astmt2 -35 -34 -30 -17 -40 -20 -30 m5astmt3 -34 -30 -39 -14 -40 -20 -30 m5astmt4 -27 -39 -12 -40 -20 -20 table 3. overview of all transition temperatures (all values in celsius); nomenclature e.g. m3isot1: material m3, iso test standard and 1 mm thick specimens. t e t rat1.2 t s m1isot1 226 = x 223 = x 224 = x m1isot1 216 = x · 0.955 214 = x · 0.959 216 = x · 0.964 m1isot3 217 = x · 0.960 217 = x · 0.973 215 = x · 0.959 m1astmt1 233 = x · 1.030 244 = x · 1.094 243 = x · 1.084 m1astmt2 226 = x · 1.000 228 = x · 1.022 223 = x · 0.995 m1astmt3 230 = x · 1.017 230 = x · 1.031 223 = x · 0.995 m1astmt4 228 = x · 1.008 228 = x · 1.022 223 = x · 0.995 m2isot1 223 = x 218 = x 223 = x m2isot2 210 = x · 0.941 207 = x · 0.949 208 = x · 0.932 m2isot3 209 = x · 0.937 207 = x · 0.954 208 = x · 0.932 m2astmt1 233 = x · 1.044 237 = x · 1.087 236 = x · 1.058 m2astmt2 218 = x · 0.977 218 = x · 1.000 213 = x · 0.955 m2astmt3 221 = x · 0.991 223 = x · 1.022 213 = x · 0.955 m2astmt4 219 = x · 0.982 213 = x · 0.955 m3isot1 254 = x 253 = x 256 = x m3isot2 245 = x · 0.964 244 = x · 0.964 243 = x · 0.949 m3isot3 247 = x · 0.972 246 = x · 0.972 248 = x · 0.968 m3astmt1 250 = x · 0.984 270 = x · 1.067 263 = x · 1.027 m3astmt2 247 = x · 0.972 259 = x · 1.023 254 = x · 0.992 m3astmt3 247 = x · 0.972 260 = x · 1.027 254 = x · 0.992 m3astmt4 249 = x · 0.980 245 = x · 0.957 table 4. determination of individual shift factors (all values in kelvin). avg. std. α 0.953 0.011 iso1mm to iso2mm β 0.958 0.015 iso1mm to iso3mm γ 1.053 0.035 iso1mm to astm1mm δ 0.993 0.022 iso1mm to astm2mm � 1.000 0.026 iso1mm to astm3mm ζ 1.005 0.038 iso1mm to astm4mm table 5. average kelvin temperature shift factors α, β, γ, δ, � and ζ and their standard deviations. the standard deviation of the shift factors listed in table 5 become greater for thicker specimens. the predictions from 1 mm iso to thicker iso series shows accurate predictions, while extrapolations from iso to astm generally hold higher deviations. figure 6 compares the predictions of the average shift factors with the results of the measured data. 69 f. kiehas, a. kalteis, m. jerabek, z. major acta polytechnica ctu proceedings (a) m1 (b) m2 (c) m3 figure 6. shift factor predictions versus reality: the trend between iso 1 mm (black) and thicker specimens (purple) can be predicted (blue) with temperature shift factors α, β, γ, δ, � and ζ. 3. conclusions the experimental investigations show that transition temperatures for elastomer modified materials hold similar trends between test standards and specimen thicknesses. temperature shift factors can be established to allow for the prediction of transition temperatures for elastomer filled pp at different testing conditions. while the data extrapolation is more consistent for iso 6603-2 than astm d3763, satisfying values for all tested materials and conditions can be achieved. this may drastically reduce testing time, effort and costs and allow for comparison of data obtained by different test standards. list of symbols fm peak load [kn] sm deflection at peak load [mm] em energy up to peak load [j] fp puncture force [kn] sp puncture deflection [mm] ep puncture energy [j] fy0 initial force [kn] sy0 initial deflection [mm] fy yield equivalent force [kn] sy yield equivalent deflection [mm] eel elastic energy portion [j] epl plastic energy portion [j] klin relative stiffness [n/mm] ∆sw width of plastic plateau [mm] rel elastic energy ratio [–] rduc ductility index [–] stotal total deflection [mm] y d yielding with deep drawing y s yielding with stable crack growth y u yielding with unstable crack growth ny no yielding t temperature [◦c] t e transition temperature; puncture energy [◦c] t rat1.2 transition temperature; elastic energy ratio [◦c] t duc30 transition temperature; ductility index [◦c] tgm transition temperature; optical [◦c] tn y transition temperature; curve type ny [◦c] ty u transition temperature; curve type yu [◦c] α,β,γ,δ,�,ζ average shift factors [–] acknowledgements this research was supported by borealis polyolefine gmbh (linz, a.) references [1] e. baur, t. a. osswald, n. rudolph. saechtling kunststoff taschenbuch. 31st printing. carl hanser verlag, 2013. [2] plastics -determination of puncture impact behaviour of rigid plastics part 2: instrumented impact testing. iso 6603-2, international organization for standardization, geneva, switzerland, 2000. [3] astm d3763-15, standard test method for high speed puncture properties of plastics using load and displacement sensors. astm international 2015. doi:10.1520/d3763-15. [4] d. huber. temperaturgeführte durchstoßprüfung von kunststoffen vergleich von durchstoßprüfmaschinen, bachelor of science, montanuniversität leoben, 2009. [5] m. böhning, u. niebergall, a. adam, w. stark. influence of biodiesel sorption on temperature-dependent impact properties of polyethylene. polymer testing 40:133 – 142, 2014. doi:10.1016/j.polymertesting.2014.09.001. [6] v. hristov, r. lach, w. grellmann. impact fracture behavior of modified polypropylene/wood fiber composites. polymer testing 23(5):581 – 589, 2004. doi:10.1016/j.polymertesting.2003.10.011. [7] a. pavan, j. williams. fracture of polymers, composites and adhesives. volume 27. elsevier science, 2000. [8] a. van der wal, a. wal, j. mulder, r. gaymans. fracture of polypropylene: 2. the effect of the crystallinity. polymer 39(22):5477–, 1998. doi:10.1016/s0032-3861(97)10279-8. [9] a. wal, a. van der wal, j. mulder, et al. polypropylenerubber blends:1. the effect of matrix properties on the impact behaviour. polymer 39(26):6781–, 1998. doi:10.1016/s0032-3861(98)00170-0. [10] a. van der wal, a. wal, r. gaymans. polypropylene-rubber blends 5: deformation mechanism during fracture. polymer 1999(40):6067–6067, 1999. doi:10.1016/s0032-3861(99)00216-5. 70 http://dx.doi.org/10.1520/d3763-15 http://dx.doi.org/10.1016/j.polymertesting.2014.09.001 http://dx.doi.org/10.1016/j.polymertesting.2003.10.011 http://dx.doi.org/10.1016/s0032-3861(97)10279-8 http://dx.doi.org/10.1016/s0032-3861(98)00170-0 http://dx.doi.org/10.1016/s0032-3861(99)00216-5 vol. 18/2018 investigation of pp-compounds for instrumented puncture tests [11] c. grein, k. bernreitner, a. hauer, et al. impact modified isotatic polypropylene with controlled rubber intrinsic viscosities: some new aspects about morphology and fracture. journal of applied polymer science 87(10):1702–1712. https://onlinelibrary. wiley.com/doi/pdf/10.1002/app.11696 doi:10.1002/app.11696. 71 https://onlinelibrary.wiley.com/doi/pdf/10.1002/app.11696 https://onlinelibrary.wiley.com/doi/pdf/10.1002/app.11696 http://dx.doi.org/10.1002/app.11696 acta polytechnica ctu proceedings 18:66–71, 2018 1 introduction 2 experimental investigation 2.1 test parameters 2.2 materials and test schedule 2.3 data evaluation 2.3.1 curve characterization 2.3.2 optical characterization 2.4 ductile/brittle transition temperature 2.4.1 puncture energy 2.4.2 elastic energy ratio 2.4.3 ductility index 2.4.4 deflection 2.4.5 curve characterization 2.4.6 optical characterization 2.5 overview transition temperatures 2.6 transition shift factors 3 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0088 acta polytechnica ctu proceedings 22:88–93, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app recycling of construction waste using high-speed milling process: determination of waste concrete zdeněk prošeka, b, ∗, pavel tesáreka, jan trejbala, tereza horováa a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague, czech republic b czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: zdenek.prosek@fsv.cvut.cz abstract. this paper deals with the use of high-speed milling process for recycling old concrete and direct determination of the potential of input waste. for this purpose, three different types of waste concrete were used: prefabricated railway sleeper, structural concrete of monolithic pillar and prefabricated drainage gutter. the paper directly examines the chemical and phase composition by xrf, x-ray diffraction (xrd), energy dispersive x-ray spectrometry (eds) and microscopic analysis, particle size distribution and ph of the recycled material. results of those analysis are used to select suitable recycled material. the suitability of choice is supported by mechanical tests of 28-day old cement pastes, where the compressive strength and dynamic modulus of elasticity are observed properties. specimens measuring 40 × 40 × 160 mm are composed of 70 wt.% portland cement and rest is micronized concrete. in all cases, the results are compared with the reference material. keywords: recycling, concrete waste, cement composites, xrd, xrf, mechanical properties. 1. introduction construction produces the largest amount of waste from all sectors, for example in the eu, construction waste is almost 33% of the 2.5 billion tons of waste generated [1]. there is an effort to recycle this waste and reduce its storage on dumps according to the eu directive. nowadays, ca. 90% of waste is recycled and the remaining 10% is stored at the dumps [2]. concrete is the most used construction material and therefore sources of concrete recyclate are relatively large and easily accessible. however, due to different properties of old concrete, the production and consequently the quality of the recyclate differ greatly. the most important aspects that affect the final quality of recyclate are: the type of use of old concrete and thus the recipe (the quality of cement and aggregate), the age of concrete and the method of care (prefabricate, monolith). there are two main areas in the construction sector, from which the concrete recyclate is obtained, the road construction and the civil engineering. the highest level of recycling is in the field of road construction because the existing road construction can be processed directly on site using mobile recycling units and reused in the construction of a new one [3, 4]. in other cases, the old concrete is taken to a landfill or recycling facility. materials from recycling lines are most often used as coarse or fine aggregates in new concrete or in stabilizing layers [5, 6]. in this way, only coarse fractions are processed. in the case of a very fine fraction (grain size below 1 mm), there is still no ideal solution to allow its use on a wider scale. several possible directions of application of this material are examined in cement composites [7–10] for example as thermally activated binder replacement [11], as mixtures for cement production [12] or as a part of geopolymers based binders [13]. due to high demands on energy consumption and co2 production the most sophisticated and environmentally friendly method is when very fine fraction of recycled concrete is used in its raw form. for this reason, the paper deals with the possibility of mechanical activation of concrete recyclate with high-speed milling. thus treated concrete may be used without other energy-intensive treatment or without auxiliary additives. mechanical activation of the recyclate occurs when hydrated cement paste is grinded and this reveal still unhydrated grains of clinker. this is primarily c2s which reacts in the long run. the amount of unhydrated clinker is directly dependent on the type of cement and the age of the concrete. in general, the amount in the cement matrix can be estimated between 10 to 20%, that 10 – 20 wt.% of recyclate can form a binder in the future composite and the remaining amount will form a filler. recyclate not only reduces the amount of cement but also the amount of natural filler in future concrete [14, 15]. 2. materials and samples the research investigated three old types of concrete that were recycled to finely ground concrete powder using high-speed milling. the first finely ground concrete recyclate was obtained from the pb2 and the sb8 railway sleepers. these sleepers had been roughly crushed first and then metal components and 88 http://dx.doi.org/10.14311/app.2019.22.0088 http://ojs.cvut.cz/ojs/index.php/app vol. 22/2019 recycling of construction waste. . . reinforcement were separated. in this way, a coarse concrete recyclate, fraction 0 – 32 mm, was obtained from which the fraction 0 – 16 mm was separated and then milled. the second finely ground recyclate was obtained from concrete drainage gutter which did not contain reinforcement and therefore the individual damaged fittings were crushed and milled. just as in the case of sleepers, this is a prefabricated product with the difference that the concrete drainage gutter is considerably younger than sleepers that are from 10 to 50 years old and are made of aggregate and cement with worse properties. the last (third) material is finely ground concrete column. these columns are from monolithic structure, where waste was generated during the reconstruction of the existing hall of the canceled factory walter motors which was founded by josef walter in 1911. the recycling process was dealt with in this material and was therefore divided into two investigated materials. the recycling process was optimized and large amount of the original filler was removed for the column a, because high-speed milling was applied only to the fraction 0 – 1 mm that originated after crushing coarse fraction of recyclate on small-volume wedge mill. the second column was the column b, which was also made of concrete recyclate fraction 0 – 16 mm. these input materials were also investigated with reference portland cement cem i 42.5r from radotín area. from these materials the samples were additionally made for testing micromechanical properties. tested materials contained 70 wt.% of portland cement cem i 42.5r and 30 wt.% of finely ground concrete recyclate. water ratio was equal to 0.35. for each recyclate a kit was created which contained five specimens with dimensions 40 × 40 × 160 mm. the second day after production, the specimens were demoulded and kept in 100% humidity, the temperature was 22 ± 1 °c during 28 days. 3. experimental methods to select the suitability of individual recyclates, a set of experimental methods were selected to describe the suitability of the material for its subsequent use. these methods include x-ray fluorescence analysis (xrf) for detecting the amount of inappropriate (foreign) substances, ph measurements for finding alkalinity of recyclates, granulometry for finding the effect of high-speed milling, xrd for determination of amount of clinker materials and electron microscopy. the suitability of selection was supported by mechanical tests of 28 days old cement pastes. x-ray fluorescence analysis (xrf) is a method based on secondary x-ray radiation and its output is a line spectrum in which the number of pulses per second is displayed for the respective wavelengths or respective energies. from this spectrum the weighting of individual elements is then calculated. elemental analysis by xrf spectrometer was performed on a spectro xepos (spectro cs s. r.) instrument. the ph was determined from aqueous extracts (in a weight ratio of 1:100) after 24 hours at room temperature (25 ± 1 °c) to ph meter (from monokrystaly a.s.). this type of ph meter is suitable for fast measurement with a deviation of about 0.2 degrees. to determine fraction of finely ground powder after the micromilling process and to determine the size of the medium grain a laser granulometer was used, the socalled dry method was used. for the determination of particle morphology and microscopic structural and elemental analysis a scanning electron microscope with the schottky cathode feg sem merlin zeiss was used, which is located in the laboratory of electron microscopy and microanalysis at the university center of energy efficient buildings. qualitative and quantitative analysis of the chemical composition of the samples was performed using an x-ray microanalysis, namely a oxford instruments energy spectrometer (eds). nondestructive resonance method was used to measure dynamic modules. the brüel & kjaer measuring set consists of the brüel & kjær type 3560-b-120, the brüel & kjær type 4519-003 acceleration sensor, the brüel & kjær type 8206 impact hammer and the control laptop. compressive strength was determined using a heckert, a model fp100 (hydraulic press). compressive strength was determined by a one-shot pressure test. testing was controlled by shifting at a constant velocity of 0.3 mm/s. 4. results and discussion table 1 summarizes x-ray fluorescence analysis results. the table lists the elements, respectively. their oxides, which were present in a concentration above 1 wt.%. the results compare individual samples of micronized recyclates with reference portland cement cem i 42.5r radotín. none of these recyclables contain excessive amounts of harmful substances. in columns a and b samples, there is a clear difference in the amount of si, which is the percentage of siliceous sand by about 8 wt.%. the presented results also showed that samples contained increased content of iron which is caused by type of milling during which the gear wheel is worn. loi means immeasurable elements such as carbon and hydrogen. the most interesting part of the results is the amount of cao because it indicates the amount of unhydrated clinker and c-s-h gel. the best properties has the column a, the drainage gutter and the railway sleeper. in another measurement, ph was measured. both recyclates and reference cement were used as samples. the resulting ph of the reference cement was 12.6. the resulting values of the recyclates themselves range from 11.5 to 12.6 ph. the results show that none of the recycled material will reduce the alkalinity of the mixture and thus negatively affect the hydration. 89 z. prošek, p. tesárek, j. trejbal, t. horová acta polytechnica ctu proceedings oxides percent weight of individual oxides [%] cement sleeper gutter column a column b sio2 17.16 40.53 66.54 36.35 44.49 cao 66.67 21.09 22.18 23.80 16.37 al2o3 3.68 11.79 22.18 7.56 10.66 fe2o3 3.43 4.10 6.37 3.13 3.64 mgo 2.35 2.29 3.93 1.42 2.03 k2o 0.67 1.83 4.52 1.45 1.64 na2o 2.97 1.51 4.32 1.35 1.84 so3 4.96 1.21 1.06 1.57 1.19 loi 3.08 15.63 3.95 23.34 18.11 table 1. results of xrf analysis. figure 1. grain curves for cement and individual samples of recyclates. figure 1 compares typical granulometric curves of reference cement and individual recycled materials. recyclates from the gutter and the column a have a finer grain curve than reference cement (radotín). in contrast, sleepers and the column b are thicker than reference cement. it can be assumed that the recyclates from the gutter and column a will be more reactive whereby we can expect higher amount of unhydrated clinker grains. figure 2 shows a record of xrd analysis with individual peaks (extremes) belonging to the respective phases. the clinker minerals c3s and c2s have basically 5 strong peaks from 32.3 to 34.4 (figure 2), of which the peak at 32.65, 33.25 and 34.4 are c3s and c2s only, or are not significantly affected other phases. peak at 32.65 labeled “c3s and c2s” is basically the only one given by the c3s and c2s phases. to calculate the representation of individual clinker materials, a peak 32.65 was used because it is basically different and is composed only of these phases. for portland cement cem i 42.5r, the amount of clinker minerals c3s and c2s was 90% ± 5%. in the case of samples gutter and sleeper, the peak at 32.65 is not visible. it means that there is not more than 1% of c3s and c2s. in case of column a, peaks are visible at 35.65, 33.25 and 34.4 but is not visible at 33.25. the results showed that the column a and column b contain clinker grains. amount of clinker is larger for the sample from column a. namely, the sample from the column a contained approx. 6.3 wt.% unhydrated clinker and the sample from column b 4.5 wt.%. table 2 shows the results from an electron microscope analysis where 100 grains were measured for each material and elemental analysis was performed for each grain and the grain size was measured. from the elemental analysis, phases were subsequently derived by stoichiometry. high speed milling produses grains in one direction, so the shape index was created, therefore the ratio of the widest part of the grain to the narrowest. results showed that the aggregate formed from quartz provides a larger grain size, moreover, grains during milling are not so unidirectional. the results of the previous five analysis showed the most suitable material for recycling the column a, the column with a larger part of the aggregate removed. to confirm this assumption, mechanical tests were carried out. results of these tests are shown in figure 3. the results of the compressive strength 90 vol. 22/2019 recycling of construction waste. . . figure 2. phases of spectra from xrd analysis. materials phase fraction average minimum grain size average maximum grain size shape index [µm] [µm] [µm] sleeper aggregate 5 210 41.4 61.3 1.5 c-s-h gel + clinker 3 250 29.1 46.5 2.0 portlandite 2 15 2.3 5.2 2.1 gutter aggregate 3 270 40.3 68.9 1.6 c-s-h gel + clinker 2 220 31.8 53.3 2.0 portlandite 7 25 9.7 15.2 1.5 column a aggregate 5 50 16.4 24.5 1.6 c-s-h gel + clinker 3 30 10.2 14.1 1.9 portlandite 5 15 5.1 10.3 2.0 column b aggregate 2 40 14.8 22.4 1.3 c-s-h gel + clinker 2 50 15.4 20.2 1.9 portlandite 5 30 10.2 30.3 1.8 table 2. results of eds analysis. 91 z. prošek, p. tesárek, j. trejbal, t. horová acta polytechnica ctu proceedings figure 3. mechanical properties of tested composites, compressive strength – left, dynamic modulus of elasticity right. and the dynamic modulus of elasticity confirm the previous assumptions and that the recyclable is best seen as column a, which reached a strength of 85.37 ± 5.68 mpa, ie by 7 mpa less than reference cement and 8 mpa more than other recyclates. 5. conclusion this work deals with the use of high speed grinding for recycling of old concrete and the direct determination of the potential of the input waste. four different materials were tested and the results show that: • none of the recyclates contained excessive amount of harmful substances, • ph of the recyclates has a similar value to the reference mixture and thanks to that none of the recyclates significantly influences ph of the hydration process, • granulometry has shown that aggregate has a major effect on the resulting granulometry because the gutter, which contained worse aggregate and the column a, where most of aggregate has been removed, have reached the required granulation for activation, • xrd analysis showed that the largest amount of unhydrated clinker and thus the bigger potential for activation had recyclate from the monolithic structure of the column, • results of the mechanical tests were confirmed by the previous analysis and that the recyclate from the column a appeared to be the most appropriate because it reached 10% stronger than the other materials. in the future, the research will focus on other concrete waste, which will be pre-recycled. acknowledgements this was financially supported by faculty of civil engineering, czech technical university in prague (sgs project sgs16/201/ohk1/3t/11) and by grant agency of the czech republic (gačr 17-06771s). references [1] eurostat. waste statistics. http://epp.eurostat.ec.europa.eu/. accsessed: 2018-09-27. [2] european parliament and council of the european union. directive 2008/98/ec of the european parliament and of the council. https://eur-lex.europa.eu/legal-content/cs/txt/ pdf/?uri=celex:32008l0098&from=lv. accessed: 2018-09-27. [3] j. r. jimenez, j. ayuso, f. agrela, et al. utilisation of unbound recycled aggregates from selected cdw in unpaved rural roads. resources, conservation and recycling 58:88–97, 2012. [4] a. pasandín, i. pérez. overview of bituminous mixtures made with recycled concrete aggregates. construction and building materials 74:151–161, 2015. [5] f. rodrigues, m. t. carvalho, l. evangelista, j. de brito. physical-chemical and mineralogical characterization of fine aggregates from construction and demolition waste recycling plants. journal of cleaner production 52:438–445, 2013. [6] j. xiao, w. li, y. fan, x. huang. an overview of study on recycled aggregate concrete in china (1996-2011). construction and building materials 31:364–383, 2012. [7] j. schoon, k. de buysser, i. van driessche, n. de belie. fines extracted from recycled concrete as alternative raw material for portland cement clinker production. cement and concrete composites 58:70–80, 2015. [8] z. shui, d. xuan, w. chen, et al. cementitious characteristics of hydrated cement paste subjected to various dehydration temperatures. construction and building materials 23:531–537, 2009. [9] y. j. kim, y. w. choi. utilization of waste concrete powder as a substitution material for cement. construction and building materials 30:500–504, 2012. [10] f. tomosawa, t. noguchi. integrated design and environmental issues in concrete technology, chap. towards completely recyclable concrete, p. 263–272. e & fn spon, london, 1996. 92 http://epp.eurostat.ec.europa.eu/ https://eur-lex.europa.eu/legal-content/cs/txt/pdf/?uri=celex:32008l0098&from=lv https://eur-lex.europa.eu/legal-content/cs/txt/pdf/?uri=celex:32008l0098&from=lv vol. 22/2019 recycling of construction waste. . . [11] z. shui, d. xuan, h. wan, b. cao. rehydration reactivity of recycled mortar from concrete waste experienced to thermal treatment. construction and building materials 22(8):1723–1729, 2008. [12] j. schoon, l. van der heyden, p. eloy, et al. waste fibrecement: an interesting alternative raw material for a sustainable portland clinker production. construction and building materials 36:391–403, 2012. [13] s. ahmari, x. ren, v. toufigh, l. zhang. production of geopolymeric binder from blended waste concrete powder and fly ash. construction and building materials 35:718–729, 2012. [14] j. topič, z. prošek, j. fládr, p. tesarek. influence of fineness of recycled concrete powder on the heat evolution and influence of its amount on mechanical-physical properties of cement paste. waste forum 2:268–274, 2018. [15] z. prošek, j. topič, j. ďureje, j. trejbal. comparison between the impact of micronized waste concrete and marble additions on mechanical properties of cementitious pastes. waste forum 2:262–267, 2018. 93 acta polytechnica ctu proceedings 22:88–93, 2019 1 introduction 2 materials and samples 3 experimental methods 4 results and discussion 5 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.11.0097 acta polytechnica ctu proceedings 11:97–102, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app analysis of dynamic effects acting on railway crossings daniela vukušičová institute of railway structures and constructions, faculty of civil engineering, brno university of technology, brno, czech republic correspondence: vukusicova.d@fce.vutbr.cz abstract. the paper is focused on measurement and analysis of dynamic effects on railway turnouts. two same-type turnouts with different fastening elasticity were chosen. attention was focused mainly on the crossing part of the turnout, where the highest dynamic impacts occur. the point of the paper is the assessment of influence of soft rail pads on spread of dynamic energy through the construction. keywords: turnout, crossing part, dynamic effects, vibrations. 1. introduction switches and crossings are the key part of a railway track. in terms of dynamic effects they belong to the most loaded constructions on the railway infrastructure especially crossing part with common crossing. they are not only the places where wheel moves from wing rail to frog (or vice versa), but also places where the stiffness changes. however the length of track with switches and crossings constitutes just a small part of the railway network, maintenance of the crossings generates relatively high cost. [1] it is primarily caused by complex forces action which is induced by passage of the train through the crossing and also by the need of maintenance of many turnout components. the study of dynamic effects will help us to clarify some regularities which may be used for future optimisation of the construction and it should lead to its service life extension. 2. aims the main aim is a comparison of dynamic behaviour of two constructions which different fastening system elasticity. the comparison is based on evaluation of transmission of vibration through these constructions. it is required to repeat measurements several times in reasonable time periods to compare the constructions in terms of increasing dynamic effects in time, as well. one of important things to do is also a choice of suitable mathematic apparatus for evaluation of dynamic effects in time, frequency and time-frequency domain. 3. stress of a crossing part the most important place of a crossing part in terms of dynamic effects is the area of transition from a wing rail to crossing nose, where the dynamic impact occurs. this impact is transmitted through the bearers to ballast which is then extremely stressed and it leads to abrasion of gravel particles on contact with a bearer. due to this stress, degradation of ballast shape under the bearer happens and it causes inadequate support of the turnout. if a turnout is not supported appropriately, the geometry of transmission forms a wing rail to crossing nose collapses and the whole degradation process accelerates. the crossing part is stressed more and more which can cause damage of a frog. 4. mechanism of wheel transition from the wing rail to the crossing nose mechanism of wheel transition from the wing rail to the crossing nose is a very complex spatial problem. the wheel passes three phases during the transition of the frog. at first, it goes on the wing rail, then it goes through a transition zone on both wing rail and crossing nose and at last it goes on crossing nose and related rail. this mechanism is shown in fig. 1 in ideal case of geometry of transition a wheel passes from the wing rail to a crossing nose continuously without dynamic impacts. this complex problem is often simplified. the simplification consists in the fact that a wheel runs on the wing rail, which deviates from the axis of the track to create a space for a crossing nose. due to the geometry of the wheel and the wing rail the wheel often descends slightly at an angle α1. after impact on the crossing nose a wheel rises again at an angle of a crossing nose α2, as long as it does not reach the starting position at the end of the transition zone. this could be seen on the simplified model of the wheel movement through the crossing part shown in fig. 2. from fig. 2 it is evident, that the total angle, at which wheels impact the crossing nose is α = α1 + α2, in this angle the influence of transition geometry on dynamic effects is simplified described. 97 http://dx.doi.org/10.14311/app.2017.11.0097 http://ojs.cvut.cz/ojs/index.php/app daniela vukušičová acta polytechnica ctu proceedings figure 1. schematic representation of the wheel set crossing over the frog – view from above [2]. figure 2. simplified model of wheel movement through the crossing part [3]. 5. measuring location and description of measured turnouts the measurements were carried out at the railway station usti nad orlici on main lines turnouts no. 3 and 4 (fig. 3). maximum speed on main line is 130 km.h−1 (160 km.h−1) for tilting vehicles, such as pendolino). both chosen constructions are the same except elasticity of the fastening system. description of the turnouts: • part of single crossover • rails profile uic 60 • fastening system (ribbed baseplate with vossloh tension clamp skl 24), turnout no. 3 has new ribbed baseplates with different rail pads – reduce overall stiffness of the track. • monoblock crossing from manganese steel • concrete sleepers • ballasted track • crossing angle 1:12 and radius 500 m • used almost exclusively in trailing direction three measuring campaigns were made on both above figure 3. turnout no. 4 and single crossover with turnout no. 3. mentioned turnouts. 6. methodology of the measurement the certified methodology of the measurement was used [4]. measurement methodology is designed for in-situ measurement in condition of full operation and was certified by the ministry of transport of the czech republic. the certified methodology consists of bearers displacements measurements, vibration acceleration measurements (on crossing, bearers and into the ballast), and pressure measurements (in two levels into railway substructure). for the purposes of this paper only small part of vibration acceleration measurement is described, only the part which is used below in evaluation part of the paper. the focus is on vibration transition from the crossing trough bearer to the ballast. the three-axis accelerometer was placed on flange of the wing rail (fig. 4) and one-axis sensor was placed on bearer nearest to the crossing nose. piezoelectric vibration acceleration sensors on wing rail and on bearer were attached to plastic handles and glued to the cleaned surface of the measured structure. 98 vol. 11/2017 analysis of dynamic effects acting on railway crossings figure 4. the three-axis accelerometer placed on flange of the wing rail. for measurement into the ballast the measuring stone was developed. the three-axis accelerometer was inserted into the measuring stone which was placed into the ballast directly under the crossing nose. 7. mathematical apparatus when dynamic action from railway traffic is measured in-situ it is necessary to describe and evaluate stochastic signal. this is rather difficult. in order to get necessary information and compare the dynamic action on each construction it is possible to evaluate stochastic signal from three different perspectives. first perspective is time domain. the calculated value is area under the curve of moving rms (rms stands for root mean square). thanks to this time perspective it’s possible to determine total dynamic action on construction. for detailed analysis of the dynamic actions above described method is not sufficient because frequency composition is unknown. for this reason it is useful to transform signal from time domain to frequency domain. for this transformation we use fourier transformation method. in frequency domain it is possible to determine which parts of signal are the strongest ones. for detailed analysis it is useful to use third perspective, which evaluate signal from time-frequency domain [5]. with this perspective we can see not only frequency composition but also its occurrence in time. 8. the welch method the welch method is certain modification of the fast fourier transformation. digital signal x[n] (n = 0, 1, 2, ...,n − 1) is divided into k segments each of them with length m (xi[m], i = 0, 1, ...,k − 1,m = 0, 1, ...,m − 1). segments are either placed in row one by one then n = k∗m or they overlap. each segment is weight by function w[m]. after transformation and following calculations periodograms components sj[k] are created. these components put together represent approximate spectral density s[k]. this estimation is described by following formulas. component of periodograms is determined by formula 1 [6]. sj[k] = 1 u ∗ m ∗ [m−1∑ m=0 x[m+i∗m]∗w[m]∗�( −j2πmk m ) ]2 (1) where u = 1 m ∗ m−1∑ m=0 w2[m] (2) is vector’s standard of window function, w[m] is window function. resultant estimation is done by averaged component periodogram. s = 1 k ∗ k−1∑ i=0 sj[k] (3) 9. short time fourier transform (stft) stft provides compromise between time and frequency signal representation [7]. its integral definition is stft (ω) x (t ′,f) = ∫ ∞ −∞ [ x(t)∗g(t−t′) ] ∗�−i2πf(t−t ′)∗dt (4) where g is window function, ’*’ complex conjunction, t′ is time displacement of window, x(t) its time representation of signal and stft (ω)x (t′,f) is its timefrequency representation [7]. 10. evaluation of vibration acceleration evaluation of vibration acceleration is divided to three sections: time domain, frequency domain and timefrequency domain. for time analysis comparable trains that passed over both turnouts were chosen. for detailed frequency analysis were chosen locomotive type 380 and two multiple units (leo express and pendolino) because they passed over both turnouts the maximum line speed. the most interesting is transmission of vibration from wing rail (a4z) through bearer (a3z) to the ballast (a33z). the evaluated signals are identified as follows: first letter a indicates vibration acceleration, number in the middle is a measuring channel number and the last letter indicates direction (z is vertical direction). 11. evaluation in time domain from many graphs and tables i chose the one which best represents results of my analysis. table 1 represents areas under the curve of moving rms on both turnouts calculated from vibration signals measured from selected trains and their comparison in all measurements campaigns. red colour typifies percentage loss of dynamic energy by transmission from rail to bearer and to ballast. value of the area under the 99 daniela vukušičová acta polytechnica ctu proceedings curve of moving rms on rail (a4z) is thought as 100% and it is calculated how many percent of this value is transmitted to bearer (a3z) and to ballast (a33z). it is obvious that values on wing rail (a4z) are increasing especially from third campaign on turnout no. 4. this fact is apparently caused by degradation of transition geometry, as the turnout no. 4 has bigger bearer displacement. on the bearer (a3z) there are also increasing values from third measuring campaign, but the increase is not significant in either of turnouts. on the turnout no. 4 there is slight increase of the percentage of dynamic energy transfer to the bearer, while on turnout no. 3 not, which is interesting. into the ballast both turnouts are comparable, but from overall point of view more dynamic energy is transferred from turnout no. 4. if we put into account all measurement campaigns on turnout no. 3, approximately 24% of dynamic energy is transmitted to the bearer, while on turnout no. 4 it is about 35%. into the ballast almost 5% of dynamic energy is transmitted from turnout no. 3 and 6,6% from turnout no. 4. the positive contribution of new fastening system on turnout no. 3 is shown in these results. 12. evaluation in frequency domain very important information can be obtained from frequency spectrum of the evaluated signal. for better comparison frequency curves were converted to numbers by calculating the areas under the curve. the area under frequency curve was divided into two frequency bands. first frequency band (150 – 600 hz) is the band where high frequency contact impact force occurs. this high frequency impact cases especially plastic deformation on the crossing nose and that is the reason why these high frequencies are not transferred into the bearer. second frequency band (0 – 150 hz) is the band where low frequency bending process occurs. these low frequencies are responsible for bearer displacements and ballast bed bending. based on the above mentioned on the wing rail both frequency bands were evaluated while on the bearer and in into the ballast only second frequency band (0 – 150 hz) was evaluated [8]. two multiple units (pendolino and leo express) and locomotive type 380 were selected for comparison of the turnouts (table 22). it is obvious that between second and third measurement campaign all values except that into the ballast are increased. lower values into the ballast can be explained by the fact that space between the bearer and ballast occurred. the bearer has worse support which resulted in bigger displacements. bigger bearer displacements are on the turnout no. 4, this is reflected in the values from the bearer (a3z). on the wing rail (a4z) especially on the frequency band 150 – 600 hz bigger values are on turnout no. 3, bud the difference between turnouts decreases. this fact can be caused by bigger wear of the turnout no. 3. difference decreases because of displacement on turnout no. 4 bearers are more significant than on turnout no. 3. based on above mentioned evaluation it is obvious that it is possible to valorise the wear of the crossing or bearer displacement only on the basis of vibration acceleration measurements. this is an important fact. 13. evaluation in time-frequency domain time-frequency method can be used for confirmation of conclusions from frequency and time analysis. the advantage of this method is that it can display time and value of frequency action in one chart. as an example i present time-frequency graph calculated for signal detected on bearer when pendolino train passed the crossing. as we can see on figure 5 there is time on horizontal axis and frequency on vertical axis. figure 5 shows the fact that the densest frequencies were detected in time when train wheel sets were passing the measuring point. maximum frequency range is from 70 to 80 hz and about 50 hz. 14. conclusion and recommendation within three measurement campaigns two turnouts with different fastening elasticity were compared. the comparison was based on measurement of dynamic effects in-situ. according to evaluation it is possible to declare that the new fastening system on turnout no. 3 appears to be perspective. it has a positive influence in reduction of dynamic effects on bearer and ballast. i would like to recommend observation of both measured turnouts and realize more measurements on them with use of the certified methodology at least once a year. if the positive influence of new fastening system confirmed, i would recommend its installation on more turnouts. references [1] a. m. zarembski. factors involved in turnout maintenance. railway track & structures, 1995. issn 0339016. [2] r. müller, b. nelain, j. nielsen, et al. description of the vibration generation mechanism of turnouts and the development of cost effective mitigation measures. 2013. report from project rivas, code: rivas_uic_wp3-3_d3_6_v02, version 2. [3] fischer, oberaigner, daves, et al. the impact of a wheel on a crossing. die stosswirkung eines rades auf das herzstück einer weiche. zev rail glasers annalen 129, 2005. issn 1618-8330. [4] j. smutný, l. pazdera, i. vukušič. hodnocení dynamických účinků na výhybky. certifikovaná metodika. ministerstvo dopravy – odbor kosmických aktivit a its, 2014. 124/2014-710-vv/1. 100 vol. 11/2017 analysis of dynamic effects acting on railway crossings area under curve of moving [m/s] measurement train a4z on crossing no. a3z on crossing no. a33z on crossing no.3 4 3 4 3 4 1 loko 380 49 43 12 13 3 324 31 6 6 2 loko 380 45 34 8 10 3 318 29 7 10 3 loko 380 64 40 11 14 4 317 35 6 8 1 leo express 123 93 31 34 4 426 36 3 5 2 leo express 121 89 27 30 5 622 33 4 6 3 leo express 152 105 32 37 4 421 35 3 4 1 pendolino 217 308 74 116 11 1334 38 5 4 2 pendolino 264 229 74 76 15 1428 33 6 6 3 pendolino 312 239 85 104 16 2327 44 5 10 average 149,7 131,1 39,4 48,1 7,1 8,124,1 34,8 4,9 6,6 table 1. area under curve of moving rms and the percentage transfer of dynamic energy from rail to sleeper and to ballast. area under curve of frequency spectrum hz * m/s2 measurement train a4z on crossing no. a3z on crossing no. a33z on crossing no. 3 4 3 4 3 4 0150hz 150600hz 0150hz 150600hz 0150hz 0150hz 0150hz 0150hz 1 loko 380 1735 6413 1340 4555 988 986 328 344 2 loko 380 1206 5854 1438 4688 621 855 314 391 3 loko 380 1620 6428 1747 4635 864 1311 199 153 1 leo express 675 3461 500 2463 502 448 72 75 2 leo express 449 2402 489 1932 293 379 66 90 3 leo express 604 3463 616 2750 421 604 47 32 1 pendolino 1018 3731 737 2718 634 816 96 135 2 pendolino 1000 3615 853 3074 642 626 132 126 3 pendolino 1085 3985 991 4008 794 1175 83 72 average 1044 4372 968 3425 640 800 149 158 table 2. area under frequency curve divided in to two frequency bands. 101 daniela vukušičová acta polytechnica ctu proceedings figure 5. time-frequency spectrogram calculated by stft method, turnout no. 3, pendolino train, third measurement campaign. [5] j. smutný, l. pazdera, m. bajer, i. moll. the analysis of dynamic effects in turnout structures by the method of quadratic time and frequency invariant transformations. proceedings of the sixth international symposium "computatuonal civil engineering 2008" iasi. societatea academica matei – teiu botez, iasi, romania, 2008. isbn 978-973-8955-41-7. [6] j. smutný. the modern method of the noise and vibration analysis applied on railway traffic. thesis of the doctoral dissertation. vut brno. isbn 80-214-0988-6. [7] j. smutný, l. pazdera. time-frequency analyses of real signals. 2003. isbn 80-86433-23-4. [8] i. vukušič. analýza dynamických účinků ve výhybkách. disertační práce. vysoké učení technické v brně, fakulta stavební, ústav železničních konstrukcí a staveb, 2005. vedoucí práce prof. ing. jaroslav smutný, ph.d. 102 acta polytechnica ctu proceedings 11:97–102, 2017 1 introduction 2 aims 3 stress of a crossing part 4 mechanism of wheel transition from the wing rail to the crossing nose 5 measuring location and description of measured turnouts 6 methodology of the measurement 7 mathematical apparatus 8 the welch method 9 short time fourier transform (stft) 10 evaluation of vibration acceleration 11 evaluation in time domain 12 evaluation in frequency domain 13 evaluation in time-frequency domain 14 conclusion and recommendation references acta polytechnica ctu proceedings doi:10.14311/app.2017.7.0017 acta polytechnica ctu proceedings 7:17–21, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app numerical discretisation of built structures assumption and reality m. gutermann∗, c. schröder university of applied sciences bremen, institute for experimental static, neustadtswall 30, bremen, germany ∗ corresponding author: marc.gutermann@hs-bremen.de abstract. the second half of the 20th century was affected by the introduction of electronic data processing. the numerical methods appeared to be capable of solving every engineering problem. however, dealing with real problems, an appropriate numerical discretization of built structures is in many cases a difficult task, since fundamental input data concerning the actual properties of the structure, material and mechanical boundary condition couldn’t always be met. typical difficulties are an incomplete documentation, unknown effects of structural faults and uncertain material properties. in this paper our experience in experimental assessment of structures is described and the outcome of experimental versus numerical results is compared in several examples. this paper summarizes finally both prospects and limitations of numerical simulations of built structures and shall show that the identified load bearing capacity is strongly bound to the quality of the numerical structural model. keywords: numerical discretisation, hybrid static, load tests, bridge, foundation, carousel. 1. introduction “a mathematical or mechanical model may be almost perfect it stays a model”. this citation from prof. opitz (tu dresden) points to the fact that many students and even some engineers believe in computational results without reflection since a computers calculation seems always correct. the truth is that the results are as good as their assumptions which lay hopefully always on the safe side. more and more existing structures have to be assessed according their serviceability and load carrying safety. the common procedure and solution strategies are shown in fig. 1. but even numerical assessment requires input data concerning actual material and structural properties as well as mechanical boundary conditions. these preconditions could not always be met so that not any calculated proof is successful. possible reasons are an incomplete documentation, unknown effects of structural faults and uncertainties in the modelling of the structural system with appropriate boundary conditions. another problem occurs if the structures have a historic design: the current technical recommendations are made for building of (new) structures and does not meet the requirements to assess historic designs. figure 1. solution strategies to asses load carrying capacity of built structures a parallel application of computational and experimental methods allows improving the numerical model and leads mostly to a much higher assessment of load carrying capacity, since the calibrated discretisation predicts the structural behaviour in a more realistic way (so called hybrid static fig. 1). on the basis of the experimental observation, the input parameters required for the analysis are obtained, e.g. to assess a railway bridge (see fig. 2). the modelling of boundary conditions as well as material parameters were calibrated using experimental results and led to a more realistic folded-plate-discretisation (see fig. 3). all loadings, recommended by guidelines, could be calculated using this model and sufficient load bearing capacity was proved. the advantage of subsequent extrapolating analysis is its totally non-destructive character. however, the disadvantage is the remaining uncertainty concerning the correctness of a numerical extrapolation. parametric studies or the consideration of large safety margins are means for handling these uncertainties. 17 http://dx.doi.org/10.14311/app.2017.7.0017 http://ojs.cvut.cz/ojs/index.php/app m. gutermann, c. schröder acta polytechnica ctu proceedings figure 2. measurements during operation figure 3. fe-model, parameters calibrated by measurements the following chapters give some examples how experimental data supplement numerical assessment and which advantages result hence. 2. case studies 2.1. masonry arch bridge “clevertorbrücke”, hanover the stone masonry arch bridge crosses the river leine in the center of hanover to transfer a multi-lane principal road. a recurring inspection revealed several open joints, where the mortar has been deeply washed out. an engineering office measured the actual shape of the three-center arch as well material properties. first numerical discretisation, taking only the arch into account as load-carrying member, showed that the bridge should have collapsed under its self-weight already. it seemed obvious that the chosen assumptions did not fit the actual load transfer so that a second calculation took backfill into account [1, 2]. this showed a higher load carrying capacity, but revealed two aspects: first, the proven load-carrying capacity was still too low for high frequented roads; second, the load carrying transfer was a theoretical assumption and could not be proven without experimental confirmation. since load carrying behaviour of backfill may be non-linear, we chose test loading with a special vehicle [3] (see fig. 4). figure 4. belfa test loading of stone masonry arch bridge, hanover the results showed, that there was an effective load transfer in the backing and that this influence stays constant even at higher load levels including partial safety factors (see tab. 1). with the experimental results it was possible to calibrate finite elements models afterwards, even simple beam-spring-models, since it was possible to choose all unknown parameters as spring stiffness or boundary conditions. in summary many influences (e.g. back filling, effective constraints, ...) supports in fact the load carrying system [4], but cannot be considered in calculations without experimental evaluation, especially if their load carrying behaviour may be non-linear at higher stress levels. 2.2. reinforced concrete foundations of wind energy converters in the first decade of this century several reinforced concrete foundations of onshore wind energy converters showed defects as cracks and spalling at concrete surface. some defects had simple causes as shrinkage or creeping but some defects were more severe and led to shut-down of sites (see fig. 5). one incorrect approach had been, that the foundation was preliminary constructed as a foundation of an immobile structure. in fact a wind energy converter is more likely a machine than an immobile construction. an experimental assessment of several affected construction showed under operation movements of several millimetres (f ≤ 2.3 mm) between foundation surface and steel built-in-part. a simple post-calculation by hand confirmed the assessed dimension, although they contain only deformation caused by elongation of concrete or steel as well as materials effects (shrinkage or creeping). poor workmanship may cause even higher deformations. our investigations showed clearly that in case of machine foundations, not only forces but also deformations should be taken into account, especially constructing connections between different materials. 18 vol. 7/2017 numerical discretisation of built structures results calculation measurement (fe-model with backing) strain deflection strain deflection action �[µm/m] f[mm] �[µm/m] f[mm] drive through of crane 19 0.33 5 0.35 (g = 48 t) test load bk 16/16 56 0.92 35 0.7 (incl. γ and crane) test load bk 30/30 97 1.57 85 1.2 (incl. γ and crane) table 1. comparison of results at arch crown (fe-model with backing vs. measurement) figure 5. double flange built-in part and observed defect mechanism 2.3. carousel, wuhan (china) after erection of a carousel the local authorities surveyed the construction by measuring strain at one of the main beams and compared these results with the submitted structural analysis of the construction. unfortunately these values differ more than 200 %. we were invited to conduct additional measurements to investigate this deviation. we measured strains at the main beam at several positions (see tab. 2) and compared the results with both simple bar model and 3d-fe-model (see fig. 6). the results showed that the upper sensors close to the weld (d1tk d3tk) measured much higher strains, caused to the influence of the nearby joint. sensor d4tk was situated out of the joints influence, removed more than half of the beam diameter (see fig. 6). as a result this sensor showed appropriate values compared to both bar and 3d-fe-model. figure 6. barand 3d-fe-model model incl. sensor positions in summary an experimental evaluation of calculations has to take into account which assumption had been made for the structural analysis. otherwise the comparison will fail without being well founded. 2.4. steel truss bridge, lomé (togo) the state togo built a steel truss bridge in the year 1960 to transport phosphates from inland to seaport (see fig. 7). in the last decades more and more steel girders showed increasing deterioration by corrosion, obviously caused by the rough environmental conditions (high humidity, salt water, phosphates). it was decided to rehabilitate the bridge by replacing top and bottom chord during train operation. fatigue had to be proven for the remaining girders, e.g. diagonals and support grid. a calculation had not given satisfactory results, so that additional information should be collected by on-site experiments and measurements including the verification of rehabilitation success. figure 7. steel truss bridge crossing the lac togo and tested fields the results were diverging: individual parts showed even higher values, but most stress levels were lower than calculated. the maximum reduction of more than 50 % was reached at the grid beams. the following proof of fatigue showed that almost all members had a sufficient remaining live time of 20 years. 19 m. gutermann, c. schröder acta polytechnica ctu proceedings sensor average measured strain interference cal. strain min [µm/m] max [µm/m] [µm/m] [µm/m] d1tk -147 129 ±40 ∼190 d2tk -196 89 ±50 ∼143 d3tk -259 68 ±50 – d4tk -95 128 ±30 ∼100 d5tk -62 56 ±25 – vb1 -30 21 ±1 26–47 table 2. relative deformation between foundation surface and steel built-in part (double flange system) figure 8. influence of discretisation: a) simple spatial framework b) detailed dispersed truss the experimental results were used afterwards to investigate the influence of calculational assumption on the stress level of truss members and consequentially on their remaining lifetime [5]. in the process the main emphasis was led on the discretisation of static system (see fig. 8) and load application. this showed that the chosen loading (single load, partial distributed load) varied the results of grid beams for several ten percent, as measured and analysed. this was actually an exceptional outcome, all other investigated parameters, as • connection of truss members as hinge or rigid joints • consideration of sheet plates between truss members • realistic discretisation of boundary conditions • joint stiffness between longitudinal and cross beams showed much lower influence, so that they can be generally neglected. in summary the investigations showed, that a more detailed spatial discretisation and especially a realistic load application can lead to much better results, even without experimental assessment. a simple analysis and discretisation of built structure is sufficient as a first approximation. 3. conclusions if we want to describe the physical reality using numerical discretisation we need additional information that may be gained by experiments, e.g. non-destructive testing, monitoring or load-tests. this may be cost and time consuming, but sometimes a little effort may have a great benefit. our experiences, documented in this paper, are summarized with the following statements: • a numerical model stays model, even if it seems perfect. concomitant experimental analysis helps to improve significantly both model and results. • experimental evaluation should meet calculation assumptions. otherwise they could not be compared. • many influences (e. g. back filling, effective constraints, ...) support in fact the load carrying system, but cannot be considered in calculations without experimental evaluation, especially if their load carrying behaviour may be non-linear at higher stress levels. • the calculated result is dependent on both system and boundary assumptions (especially in cases of arches and fatigue proof). change of system parameters may vary the results by 100 % or more. • discretisation of machine foundation should take (tolerable) deformations into account. an analysis as static and stiff system may not be sufficient and may lead to severe defects. acknowledgements we want to send special thanks to the university of applied sciences bremen and all involved parties, which enables us to have done more than 300 successful investigations in the field of analysing load carrying safety of built structures. references [1] d. proske, p. lieberwirth, p. h. van gelder. sicherheitsbeurteilung historischer steinbogenbrücken. mauerwerk 11(4):186–189, 2007. doi:10.1002/dama.200700327. [2] z. orbán, m. gutermann. assessment of masonry arch railway bridges using non-destructive in-situ testing methods. engineering structures 31(10):2287–2298, 2009. doi:dx.doi.org/10.1016/j.engstruct.2009.04.008. [3] m. gutermann, c. schröder. loading vehicle belfa development and experience gained in 10 years of practice. bridge structures 11(1,2):19–31, 2015. doi:10.3233/brs-150081. [4] m. gutermann. an article on experimental assessment of structural safety of solid bridges. proceedings of icem12-12th international conference on experimental mechanics 28, 2004. 20 http://dx.doi.org/10.1002/dama.200700327 http://dx.doi.org/dx.doi.org/10.1016/j.engstruct.2009.04.008 http://dx.doi.org/10.3233/brs-150081 vol. 7/2017 numerical discretisation of built structures [5] w. malgut, k. ammermann, t. sander, m. gutermann. pont lagunaire, togo experimentelle tragwerksanalyse einer stahlfachwerkbrücke zum nachweis der restnutzungszeit. 8 symposium "experimentelle untersuchungen von baukonstruktionen" 2015. 21 acta polytechnica ctu proceedings 7:17–21, 2017 1 introduction 2 case studies 2.1 masonry arch bridge "clevertorbrücke", hanover 2.2 reinforced concrete foundations of wind energy converters 2.3 carousel, wuhan (china) 2.4 steel truss bridge, lomé (togo) 3 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.14.0008 acta polytechnica ctu proceedings 14:8–13, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app uncertainty propagation for lwr burnup benchmark using sampling based code scale/sampler daniel campolinaa, b, ∗, jan fryborta a department of nuclear reactors, czech technical university in prague, prague, czech republic b nuclear technology development center, belo horizonte, brazil ∗ corresponding author: campolina@cdtn.br abstract. sampling based method is adopted in many fields of engineering and it is currently used to propagate uncertainties from physical parameters and from nuclear data, to integral indicators of nuclear systems. the total uncertainty associated with a model simulation is of major importance for safety analysis and to guide vendors about acceptable tolerance limits for nuclear installations parts. this work presents some calculations to propagate uncertainties for a nuclear reactor fuel element modeled in scale/triton, using the sampling tool scale/sampler. results showed that the influence of input uncertainties on kinf is more pronounced in the fresh core other than the depleted core and the contribution from studied manufacturing uncertainties is smaller than the contribution of nuclear data uncertainties. keywords: random sampling, uncertainty, nuclear fuel burnup, oecd uam lwr. 1. introduction in a global picture, uncertainty quantification (uq) is the process of characterizing input uncertainties, forward propagating these uncertainties through a computational model, and performing statistical assessments on the resulting responses. in this forward propagation, illustrated in fig.1, probabilistic or interval information on parametric inputs are mapped through the computational model to assess statistics or intervals on outputs. the computational model for the present application is the triton control module of scale, composed of cross-section processing codes, a neutron transport solver, and point depletion code [1]. for uq, components of the input parameter are considered to be uncertain as specified by probability distributions (e.g., normal, beta). by stipulating specific distributional structure to the inputs, distributional structure for the outputs (i.e, response statistics) can be inferred [2]. uncertainties and correlations among experimentally measured cross-sections constitute the so called nuclear data covariance libraries. the sampling based approach can be used to sample physical parameters like dimensions and densities and also joint probability density functions given in the nuclear data covariance libraries. the last produces a random sample for the nuclear cross-sections that are used in a transport calculation. one of the most important results of a nuclear reactor simulation is neutron multiplication factor (k) that measures the balance between production of neutrons and their absorption in the core and leakage out of the core. the multiplication factor is a key figure 1. realizations from specified input distributions being run in a model simulation. indicator for the changes in the neutron flux that also causes changes in the power level of the reactor. if the multiplication factor is lower than 1, the system is subcritical and its power is decreasing. loss of neutrons is greater than their production and it is impossible to maintain the fission chain reaction. a uq analysis can provide the level of confidence in the multiplication factor predicted by simulations and identify through a sensitivity analysis the key parameters whose uncertainties contribute most to the multiplication factor uncertainty. there are always safety margins to multiplication factor in criticality calculations. it reflects the uncertainty in all calculations. when subcriticality of some system must be certified the calculations are usually aimed at maximum multiplication factor by 5 % below the criticality. the uq process can verify that such a margin is sufficient for any design systems. since sampling based method only requires forward 8 http://dx.doi.org/10.14311/app.2018.14.0008 http://ojs.cvut.cz/ojs/index.php/app vol. 14/2018 uncertainty propagation for lwr burnup benchmark calculations, it presents some advantages if compared to adjoint-based perturbation theory approach used by e.g. tsunami modules of scale. in problems with significant second order effects, the first order perturbation theory is not valid while the sampling method is fully applicable. the typical example of the second order effect is when the perturbation in the transport operator causes significant perturbation in the flux solution of the boltzmann transport equation (see chapter 6 of [1]). scale/sampler is one of the sampling tools available at department of nuclear reactors of cvut for uncertainty qualification. it is at first used in this study to assess the impact of covariance data on nuclear transport and depletion of a lwr benchmark 2d model. then effect of geometric uncertainties, also known as manufacturing uncertainties, are quantified for the same model. 2. uncertainty analysis from nuclear data in this section, cross-section perturbation using scale/sampler is performed for a benchmark proposed by the oecd expert group uncertainty analysis in modeling (uam) [3]. resulted values of propagated uncertainties to kinf are compared against participant-averaged results for the uam benchmark. 2.1. methodology the uam benchmark has several exercises. this analysis is using fuel pin model of a typical pwr fuel rod in 15x15 assembly design from exercise i-b. this was proposed in scope of the uam benchmark in order to address uncertainties in nuclear fuel depletion calculation due to the basic nuclear data as well as the impact of processing the nuclear and covariance data. summary and discussion of neutronic cases performed by the participants is available in [4]. participantaveraged results for exercise i-b are reproduced here for comparison purposes despite different input covariance cross-sections were used by the participants. the uam expert group selected 44-group scale 6.0 covariance library as a reference library for the benchmark exercise since it was the most complete and up-to-date compilation at the time of that review. calculations in scale are conducted according to control modules and sequences. this study adopted t-depl sequence of triton that couples 2d multigroup deterministic transport in newt with origen module depletion. the lattice physics calculations are based on the 56-group energy structure of the v7-56 scale library. this library is recommended for light-water reactor calculations. trace amount of 231 nuclides was added in the inventory of burnable materials (addnux=3 card of triton) in order to reach more realistic fuel depletion from the very beginning. infinite lattice of the elementary cell was simulated by reflective boundary condition adopted for all sides of the model. parameter value hfp conditions fuel temperature (k) 900 cladding temperature (k) 600 moderator (coolant) temperature (k) 562 moderator (coolant) density (g/cm3) 0.7484 configuration unit cell pitch (mm) 14.427 fuel pellet diameter (mm) 9.391 fuel pellet material uo2 fuel density (g/cm3) 10.283 fuel enrichment (w/o) 4.85 cladding outside diameter (mm) 10.928 cladding thickness (mm) 0.673 cladding material zircaloy-4 cladding density (g/cm3) 6.55 gap material he moderator material h2o operating history operating cycle 1 burn time (days) 1825 final burn-up (gwd/mtu) 61.28 specific power (kw/kgu) 33.58 table 1. configuration and conditions of burnup pin-cell test problem. materials: pitch of unit cell reflective boundary r e fl e ct iv e b o u n d a ry r e fl e ct iv e b o u n d a ry figure 2. triton model model for exercise i-1b pin cell. the scale/sampler module was used to perform uncertainty propagation/analysis of neutron multiplication factor for the pin-cell shown in fig. 2. data library for cross-section perturbations is described in the following section. cross section covariance library the new 56-group (56groupcov) scale 6.2 covariance library [1], recommended by scale 6.2.1 for 9 daniel campolina, jan frybort acta polytechnica ctu proceedings kinf 0 gwd/mtu 60 gwd/mtu sampler sampler ref.* ∆ mean val 1.4023 0.9063 0.9030 0.4 % rsd (%) 0.55 0.58 0.89 sd (pcm) 770 529 805 52 % table 2. initial condition and final burnup uncertainties in kinf for cases uam* (using 44groupcov) and sampler (using 56groupcov). all applications, was used by sampler uncertainty module. the data correspond to relative uncertainties assembled from a variety of sources, including highfidelity covariance evaluations from endf/b-vii.1 [5] as well as approximate uncertainties obtained from the collaborative projects involving bnl, lanl, and ornl. the assumptions in generating the data, the library contents, and processing procedure for the scale 56-group covariance libraries can be found in [1]. comparison between the two libraries adopted by sampler and adopted by uam participants results are provided in next section in terms of relative standard deviation for some reactions, in order to clarify the difference and the probable impact on results. the number of samples/realizations in sampler code was set to 1000, that is the maximum allowed for cross section sampling. recent works show that 1000 calculations are enough to permit optimal convergence of the statistical moments of significance for any given response function [6–8]. it is important to emphasize that only cross-section (xs) perturbation is performed and it is the most important contributor to the overall uncertainty. neither fission yield perturbation nor decay data perturbations are performed. some studies considering the same benchmark concluded that contribution of decay data uncertainty for burnup calculations is negligible [9]. on the other hand, other studies stated that fission yield and decay data uncertainties may be relevant for uncertainties of nuclide densities in depletion calculations [10]. 2.2. results propagation of uncertainties to initial and final burnup results performed using sampler are presented in tab. 2 together with the mean value of kinf calculated. participant-averaged result from exercise i-b benchmark (only available for the final burnup condition in [4]) is presented in the fifth column for comparison purposes. the symbol ∆ is the difference between the value calculated in this study by sampler and the reference value. note that sd stands for 1 standard deviation and rsd is the relative standard deviation. the difference between kinf mean value calculated by sampler and the reference value was 0.4 %. it is typical difference that can be observed for criticality calculations of identical systems by different codes and xs libraries [11]. this multiplication factor value is the average for the all the 1000 perturbed cases. as the xs perturbations are based on normal distribution, figure 3. running kinf average and standard deviation as function of sample size for simulation on sampler. the mean value should stay almost the same as for a single case with average xs values. in order to provide information on sampling convergence, fig. 3 presents the average kinf and the running standard deviation as function of the sample size. it is possible to see that after 800 samples, the average of calculated kinf mean values is stabilized. in contrast to the above results, the increase of propagated uncertainty in kinf for burnup 60 gwd/mtu was 52 %. sampler estimated 529 pcm for the standard deviation on kinf and the reference value is 805 pcm. the large difference (scale x uam reference) can be explained by use of different versions of the covariance libraries. there are also sampler results for the fresh fuel and the final burnup. the estimated standard deviation on kinf was 770 pcm for the fresh fuel and it decreased to 529 pcm for the final burnup. the 31 % difference comes from change of fuel composition during depletion. each nuclide has its own xs covariance data thus the total uncertainty changes during depletion. the reason for the uncertainty decrease is discussed below and the main contributors to this phenomenon are illustrated in fig. 4 and 5. it is possible to select the main contributors to the overall kinf uncertainty. these examples can also serve as the explanation for differences between the current and the older version of the xs covariance library in scale; and they are also the reason for computed uncertainty decrease during fuel depletion. the usual main contributor to the calculation uncer10 vol. 14/2018 uncertainty propagation for lwr burnup benchmark variable distribution value std.dev lower cutoff upper cutoff fuel_radius normal 0.46955 2.1667e-4 0.4689 0.4702 gap_thickness normal 0.00955 1.3333e-3 0.00555 0.01355 clad_thickness normal 0.0673 1.3333e-3 0.0633 0.0713 all dimensions in cm. table 3. parameters for variable sampling in sampler input file. figure 4. 235u nubar relative standard deviation by energy for covariance libraries 44groupcov and 56groupcov. tainty is average fission neutron yield (nubar) of the major fissile nuclides, i.e. 235u and 239pu. fig. 4 and fig. 5 shows the relative standard deviation of this fission neutron yield in 44-groupcov and 56-groupcov for 235u and 239pu, respectively. the relative standard deviation of 235u nubar below around 10 ev is approximately by 24 % higher for 56-groupcov. much larger difference can be observed for 239p u nubar. its rsd is about three times larger in 44-groupcov when compared to 56-groupcov. it should be also noted that its uncertainty value is lower when compared to 235u nubar. it is the reason for multiplication factor calculation uncertainty decrease during fuel depletion when adopting 56-groupcov. in the final figure 6, this comparison shows that the relative standard deviation of absorption in 238u is comparable for these two covariance libraries and it is below 5 % for all neutrons except fast neutrons. in addition to the different covariance data used for the two methodologies compared, uncertainties may be introduced into the broad-group cross-sections due to approximations in the grouping procedure. in [12], it is showed that the dominant uncertainty is generally with regard to the energy weighting function used to average the point-wise data within a broader groups. figure 5. 239pu nubar relative standard deviation by energy for reproduced from covariance libraries 44groupcov and 56groupcov. figure 6. 238u(n,γ) relative standard deviation by energy for covariance libraries 44groupcov and 56groupcov. 3. uncertainty analysis from geometry perturbation the uam benchmark model of pwr fuel rod presented in section 2.1 had its geometrical uncertain parameters of fuel radius, gap thickness, and clad thickness modeled in sampler. the methodology adopted and resulting propagated uncertainty to the neutron multiplication factor are presented. 11 daniel campolina, jan frybort acta polytechnica ctu proceedings 0 gwd/mtu 60 gwd/mtu name variable # samples kinf rsd(%) sd(pcm) kinf rsd(%) sd(pcm) case1 fuel_radius 300 1.40749 0.010 14 0.90494 0.002 2 case2 gap_thickness 300 1.40749 0.004 6 0.90494 0.000 0 case3 clad_thickness 300 1.40749 0.005 8 0.90494 0.001 1 case4 all together 300 1.40749 0.012 17 0.90494 0.003 3 table 4. variables sampled for each study case and results of uq with the contributors to uncertainty in kinf. 3.1. uncertain variables normal distribution of the geometrical parameters was assumed. the uam benchmark specifies values of standard deviations and there is also requirement to truncate the distribution at 3 standard deviations as lower cutoff and upper cutoff. table 3 presents the parameters considered in the sampling of variables. uq for each variable was calculated in order to perform the sensitivity study of the geometric parameters. the values of fuel radius (fuelr), gap outer radius (gapr) and cladding outer radius (cladr) as function of the sampled variables (tab. 3) are the following: • case1 (variable sampled: fuel_radius ) fuelr = fuel_radius gapr = fuel_radius+0.00955 cladr = fuel_radius + 0.00955 + 0.0673 • case2 (variable sampled: gap_thickness) fuelr = 0.46955 gapr = 0.46955 + gap_thickness cladr = 0.46955 + gap_thickness+0.0673 • case3 (variable sampled: clad_thickness) fuelr = 0.46955 gapr = 0.4791 cladr = 0.4791 + clad_thickness • case4 (variable sampled: all together) fuelr = fuel_radius gapr = fuel_radius + gap_thickness cladr = fuel_radius + gap_thickness + clad_thickness since the cladding outer radius (cladr) depends on uncertain variables values, the moderator/fuel volume ratio and the cladding density vary according to the sampled values. however, for being a neutron nonabsorbing material (zircaloy-4), the density change was disregarded. the fuel meat density (d), dependent on fuel radius for case 1 and case 4, is updated according to equation 1. this allows to keep the value of fuel mass constant when fuel radius sampled value is updated in the model geometry. the nominal density is worth 10.283 g/cm3 according to tab. 1. dnew[g/cm3] = 10.283 × ( 0.46955 f uel_radius )2 (1) 3.2. results the cases considered are presented in tab. 4, where the variable under study is given in the second column. the contribution to uncertainty in kinf are presented for 0 gwd/mtu and 60 gwd/mtu burnups. uncertainty in the fuel radius (case 1) is the biggest contributor to the overall calculation uncertainty due to the fact that its value directly influences fuel volume. the second in rank is the clad thickness represented by case 3. it was noticed that the effect of uncertainties in kinf is much more pronounced for the fresh fuel rather than for the final burnup condition. 4. conclusions uncertainty propagation analysis was performed for the burnup pin-cell exercise i-b of the uam benchmarks. the sampler code using scale 6.2 56groupcov covariance library had the results compared against participant-averaged result from the benchmark exercise i-b. the difference between propagated uncertainty (sd) in kinf calculated by sampler and the reference case (uam benchmark) was 52 %. such difference is expected since the codes used different versions of covariance libraries with different group structure. comparison of results for the fresh and spent fuel for the same covariance library was done for the sampler results alone, as there are no benchmark reference results. the observed 31 % difference in standard deviations is attributed to the fact that propagated uncertainty is mainly affected by nuclide composition that changes along the burnup. future research must include other nuclear libraries for comparison. when propagating geometrical uncertainties, the effect of uncertainty in fuel radius, clad and gap thickness was calculated. it was noted that the influence of geometric input uncertainties on kinf is more pronounced for the fresh than for the depleted fuel. this can be explained by the fact that fresh fuel contains higher density of fissile material and influences more the neutron population of the reactor. the sensitivity study also confirmed that uncertainty of fuel radius (case 1) is the biggest contributor to the overall calculation uncertainty since it influences directly fuel volume even though the total fuel mass is preserved. this study also confirmed that the contribution of manufacturing geometrical uncertainties to the prediction 12 vol. 14/2018 uncertainty propagation for lwr burnup benchmark of kinf is smaller than the contribution of cross-section uncertainties. acknowledgements the mobility of the brazilian co-author to department of nuclear reactors, czech technical university in prague has been possible with the erasmus mundus smart2 support (project reference: 552042-em-1-2014-1-fr-era mundus-ema2) coordinated by centrale supelec. 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[2] a. haldar, s. mahadevan. probability, reliability, and statistical methods in engineering design. john wiley, 2000. [3] k. ivanov, m. avramova, s. kamerow, et al. benchmark for uncertainty analysis in modeling (uam) for design, operation and safety analysis of lwrs, vol. 1. citeseer, 2013. [4] r. n. bratton, m. avramova, k. ivanov. oecd/nea benchmark for uncertainty analysis in modeling (uam) for lwrs–summary and discussion of neutronics cases (phase i). nuclear engineering and technology 46(3):313–342, 2014. doi:https://doi.org/10.5516/net.01.2014.710. [5] m. chadwick, m. herman, p. obložinskỳ, et al. endf/b-vii. 1 nuclear data for science and technology: cross sections, covariances, fission product yields and decay data. nuclear data sheets 112(12):2887–2996, 2011. doi:https://doi.org/10.1016/j.nds.2011.11.002. 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[12] z. ting. sampling-based nuclear data uncertainty quantification for continuous energy monte carlo codes. ph.d. thesis, école polytechnique féeérale de lausanne, 2015. doi:10.5075/epfl-thesis-6598. 13 http://dx.doi.org/https://doi.org/10.5516/net.01.2014.710 http://dx.doi.org/https://doi.org/10.1016/j.nds.2011.11.002 http://dx.doi.org/https://doi.org/10.1016/j.anucene.2015.03.046 http://dx.doi.org/http://dx.doi.org/10.1155/2013/549793 http://dx.doi.org/http://dx.doi.org/10.1155/2012/767096 http://dx.doi.org/http://dx.doi.org/10.1155/2013/790206 http://dx.doi.org/https://doi.org/10.1016/j.anucene.2016.11.025 http://dx.doi.org/https://doi.org/10.1016/j.anucene.2011.01.032 http://dx.doi.org/10.5075/epfl-thesis-6598 acta polytechnica ctu proceedings 14:36–41, 2018 1 introduction 2 uncertainty analysis from nuclear data 2.1 methodology 2.2 results 3 uncertainty analysis from geometry perturbation 3.1 uncertain variables 3.2 results 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.19.0046 acta polytechnica ctu proceedings 19:46–52, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app selection of reference value of longitudinal to transversal ultrasonic waves velocity ratio on pressure purpose steels for detection of creep tomáš zavadila, b a advanced technology group, s.r.o., toužimská 771, 199 02 prague, czech republic b department of materials, faculty of nuclear sciences and physical engineering, czech technical university in prague, trojanova 13, 120 01 prague, czech republic correspondence: zavadilt@atg.cz abstract. common degradation process for industrial equipment in power and petrochemical industry is creep. the ultrasonic waves velocities and their ratio can provide feedback on stress-strain conditions generated by heat. selection of potentially critically affected parts of pressure equipment can be theoretically accessed by their means. selection of such critical parts or areas is done by comparison with reference benchmark value. setting the standard reference value for conventional pressure purpose steel is crucial so it can be later used as a benchmark for selection of critical areas on equipment under operation. the article proposes a benchmark value to be utilized for p265gh pressure purpose steel and steels of similar grade. keywords: residual stress, non-destructive testing, ultrasonic testing, creep. 1. introduction creep is one of the main phenomena responsible for the failure of plant pressure equipment operating at elevated temperatures. creep is associated with the synergistic effect of changes of microstructure and strain accumulation, which leads to nucleation, growth and coalescence of microvoids, and subsequently to component failure. extending the service life of major components is crucial in high temperature plant applications. the life expectancy is based on the ability of the material to retain its high-temperature creep strength for period of at least twice the projected design life. development of new methods that shall be able to assess the material condition is therefore necessary [1]. various non-destructive testing techniques have been developed to detect the creep damage. techniques range from acoustic emission, infrared thermography, ultrasonic attenuation and velocity measurements, acoustic harmonic measurements up to eddy currents [2, 3]. acousto-elastic and acousto-plastic effects are responding to stress in elastically deformed body and to residual stresses created by macroscopic deformation. most of residual stresses are induced by plastic strain. creep is a long-time plastic deformation of equipment subject to heat and pressure, therefore acoustic measurements of residual stresses are suitable candidate for the proposed application for detection of creep degradation process. in order to measure the response to creep degradation process it is necessary to measure the stress values itself, or compare individual measurements suspected from creep to a reference benchmark value for unaffected material. the following article is dealing with setting of this reference benchmark value. 2. stress measurement by conventional ultrasonic testing the ultrasonic waves velocity is a function of stress σ as described by obraz [4]. example of a simple uniaxial stress measurement in a bar may be written as follows: σ = δt− (βt + αt )∆t βσ + 1� (1) where δt = ∆t/t0 is fractional change in time of flight (tof), αt and βt are coefficients of heat expansion, ∆t is change of temperature and βσ is acoustoelastic coefficient and � is the fractional change in length. if the ∆t → 0 (there is no significant temperature change), the formula can be simplified as: σ = δt βσ + 1� (2) we may use the tof of either longitudinal ultrasonic waves or transversal waves. putting the tof values to a ratio, the situation further transforms to: σ = t0t t0l − tt tl (βσl + 1� )[ t0t t0l + tt tl ] (3) where t0t/t0l is tof ratio on tested part before applied loading and tt/tl after the applied loading. this formula, however, is difficult to use during inservice inspections for the following reasons: 46 http://dx.doi.org/10.14311/app.2018.19.0046 http://ojs.cvut.cz/ojs/index.php/app vol. 19/2018 selection of ultrasonic waves velocities for detection of creep (1.) the equipment’s original state is not available for comparison. (2.) the fractional change of length � cannot be calculated as the precise dimensions may not be known. (3.) the stress in operated equipment is rather multiaxial than uniaxial, especially when considering creep-induced strengthening. therefore despite the ultrasonic waves velocity (and respective tof) are changed due to the induced residual stress, it cannot be measured directly. this implies that precise measurement of stress σ cannot be used and ordinal means of evaluation (i.e. comparison of tested area with (a set of) reference samples) for detection of critical areas is the only option. the ultrasonic waves velocities for generally anisotropic material of the tested part are given (as proposed by according to schneder and goebbels [5]) as follows: czz = √ λ + 2µ + (2l + λ)ξ + (4m + 4λ + 10µ)�z ρ czx = √ µ + (λ + m)ξ + 4µ�z + 2µ�x − 0.5n�y ρ czy = √ µ + (λ + m)ξ + 4µ�z + 2µ�y − 0.5n�x ρ (4) where cij is ultrasonic waves velocity in plain ij, ρ is density, �i is strain in the given axis, ξ = �x + �y + �z, λ and µ are lamé’s constants, and l, m, n are murnaghan’s constants. young modulus e and shear modulus g are linked with lamé’s constants as described in the formula below: e = ν(3λ + 2µ) λ + µ ; g = µ (5) where ν is poisson’s ratio. assuming the tested volume of the tested part has isotropic behavior (i.e. the material properties are not changing with directions) and using the abovementioned formula (5), the equations for longitudinal and transversal waves velocity (4) may be then further simplified to the following: cl = √ e ρ 1 −ν (1 + ν)(1 − 2ν) ct = √ g ρ = √ e ρ 1 2(1 + ν) (6) where e is young modulus, g is shear modulus, ρ is density and ν is poisson’s ratio. the common observation is that the longitudinal waves velocity cl increases with applied load in elastic region while the transversal waves velocity decreases. as was experimentally confirmed by the team from det norske veritas [6], the increase of velocity of the longitudinal waves is limited and when it reaches a critical level of strain, it starts to decrease again. similar result was observed also for transversal waves by kobayashi [7] where the velocity decreases in elastic region, while in the plastic region it again increases. the behavior was described as being a consequence of anisotropy in elastic properties and inhomogeneous localization of plastic strain [8, 9]. such may include creation of point defects, slip bands and yield vertex which is responsible for degradation of elastic modulus. the inverse behavior of longitudinal and transversal waves’ velocity may be used advantageously when observed as a ratio of both values, i.e. cl/ct (also l/t ratio, dimensionless). the problem can be then transformed to a single-variable equation (only variable is the poisson’s ratio) with monotone-increasing features: cl ct = √ e ρ 1 −ν (1 + ν)(1 − 2ν) √ ρ g = √ 2(1 −ν) 1 − 2ν = f(ν) (7) this can help to determine the current stress/strain relation as was observed by kumar et al. [10] in their article about correlation of transversal (and longitudinal) wave velocity and poisson’s ratio. young and shear modulus tend to vary in the same direction in any metal material. this implies that the variation of the poisson’s ratio is dependent only on how much each of the moduli will be affected. let’s assume the range of poisson’s ratio is practically bound to 0 < ν < 0.5 (for isotropic solids) [11]. the higher the ratio the less volume change during deformation, where ν = 0.5 means no volume change. plot of the abovementioned part of the equation (7) is on fig. 1. poisson’s ratio as ν = 0.25 ÷ 0.30 is a common value for conventional structural steels and similar, so this provides preliminary expected range of l/t ratio values of cl/ct = 1.73 ÷ 1.87. 3. effect of creep on l/t ratio values ongoing research of the author [13] performed measurements on collapsed membrane wall from p265gh pressure purpose steel that was subject to creep rupture. the observed results concluded the it was shown that the l/t ratio is sensitive to plastic deformation due to strengthening mechanisms and is able to detect areas exposed to tension or compression due to bending as well as plastic deformation due to combined effect of internal pressure and heat. the values of l/t ratio in this experiment proven to be in range of 1.80 ÷ 1.85 in measurement spots far from the area affected by creep. the values were growing from the reference value towards its maximum of 2.086 located on the crack edge from all directions. the affected area of elevated l/t ratio values was < 50 mm from the crack edge. the key results are shown in fig. 2. from the fig. 2 it can be assumed that the 1.85 might be the benchmark reference value for p265gh 47 tomáš zavadil acta polytechnica ctu proceedings figure 1. cl/ct as a function of the poisson’s ratio ν for isotropic continuum. the selected area of values common for structural steels (ν = 0.25 ÷ 0.30) provides expected range of cl/ct = 1.73 ÷ 1.87 [12]. figure 2. sample subject to creep rupture with measured values of l/t ratio along the planes r1 (blue), r2 (red) and r3 (green). connecting lines in the charts do not provide any physical relation or a trend, they are for better orientation only [13]. grade pressure purpose steel. if the premise is accepted, the elevated values of 2.086 were above the 3s (3 standard deviations) of the assumed benchmark value and therefore the elevated values should be considered as relevant and referring about non-standard stress-strain conditions, potentially caused by a degradation process as creep (the type degradation process is assumed due to knowledge of the history of operation of the particular tested part). the following part of the article focuses on finding a standard benchmark value for unaffected material (i.e. non-degraded) and therefore proving the abovementioned assumption is justifiable. 4. experimental setup the experiment intends to find expected value that can be used as a standard global benchmark value of the l/t ratio measured on pressure purpose steel that is not subject to long-time processes causing high residual stresses. a set of steels with similar chemical composition (to each other as well as to the original p265gh steel) and different levels of heat treatment have been tested and mutually compared. the each of the following: (1.) expected value and standard deviation of each of the steels, (2.) value trend based on chemical composition and heat treatment, (3.) and range of values for the whole steel group, shall then help determining what are the benchmark l/t ratio values as well as what are the values that might be already considered as caused by severe degradation processes. if the values will be in controlled range for unaffected steel (i.e. standard deviation of l/t ratio values for samples with the same carbon content but various heat treatment shall be low), it can be assumed 48 vol. 19/2018 selection of ultrasonic waves velocities for detection of creep that the l/t ratio has materialand heat-treatmentspecific values. that would imply a significant change of material properties due to long exposure to heat is measurable, and clearly distinguishable from unaffected material. for this experiment, the level of l/t ratio considered as out of standard value range shall be defined as: ∆ cl ct = ∆ cl ct (σ) > 3s; σ � σ0 (8) or more precisely: ∆ cl ct (σ) > 3 ·maxj{sj} = = 3·maxj √√√√ 1 n − 1 nj∑ i=0 ( cl cti −〈 cl ct 〉j)2; σ � σ0 (9) where sj is the standard deviation of the steel j, nj is the number of measurements for the benchmark value of steel j, 〈cl/ct〉 is the average mean value of the all steel types, σ0 a generally unknown stress level of steel that is not subject to longtime exposure to heat and σ is a stress of affected (exposed, degraded) volume. the result shall be compared with the expected result acc. to fig. 1, where it should fit the expected range of values 1.73 ÷ 1.87. 4.1. samples samples subject to the experiment were short bars with a length of 20 mm and various diameter up to 20 mm. samples were from three types of standard structural steel (c16e/1.1148, c35/1.0501, c55/1.0535) heat treated in a laboratory conditions for 2 quenching temperatures tq (850 °c and 950 °c) and 4 various tempering temperatures tt (400 °c, 500 °c, 600 °c, and 700 °c) for one hour. the samples were in totally 24 sets and each set of samples had between 4 and 5 samples. the selected steel types were similar in the chemical composition to the conventional p265gh pressure purpose steel [12]. the closest steel (by chemical composition) to p265gh was c16e, other steels had mainly higher carbon content, which assumes different (higher) standard values for the l/t ratio. the heat-treatment processings were selected to simulate maximal structural changes that are not caused by extreme degradation processes (i.e. mainly phase transform and recrystallization effects). 4.2. equipment measurement was performed on sonatest sitescan 250s by ultrasonic pulse-echo technique. the standard ultrasonic testing (ut) equipment was selected on purpose to 1) simulate standard in-service inspection options of measurement and 2) to fit the measurement with analysis done on collapsed membrane wall in experiment briefly described in chapter 3. the time of flight (tof) was measured directly on the device and it was recalculated to ultrasonic waves’ velocity. conventional direct ut probes sonatest rdt2550 with peak frequency 4.64 mhz for longitudinal waves and panametrics v155 with peak frequency of 4.25 mhz for transversal waves were used. measurement areas were flat, with smooth surface and without any corrosion products that may affect the measurement. the precise thickness was measured by a micrometer. 5. results general the measured values were well controlled with l/t ratio in the range of cl/ct = 1.812 ÷ 1.835. the average l/t ratio was 〈cl/ct〉 = 1.82. the maximal nominal value as well as standard deviation s for the measurements was maxj{sj} = 0.005 for j equal to steel c55 quenched on tq = 950 °c, which was fully phase transformed to martensitic structure. steel c16e steel c16e/1.1148 is by chemical composition closest to the p265gh steel. compared to the other groups it has the lowest amount of carbon c and higher average of manganese mn. values of the l/t ratio for this steel have the average of 〈cl/ct〉c16e|850 = 1.816 ± 0.001 for group quenched on tq = 850 °c and 〈cl/ct〉c16e|950 = 1.819±0.001 for group quenched on tq = 950 °c. the values are virtually almost constant with changing tempering temperature tt . steel c35 steel c35/1.0501 has the average amount of carbon c and lower average of manganese mn. values of the l/t ratio have the average of 〈cl/ct〉35|850 = 1.816 ± 0.002 for group quenched on tq = 850 °c and 〈cl/ct〉35|950 = 1.819±0.003 for group quenched on tq = 950 °c. the values have monotonously decreasing trend with increasing tempering temperature tt . steel c55 steel c55/1.0535 compared to the other groups has the maximal amount of carbon c and lower average of manganese mn. values of the l/t ratio have the average of 〈cl/ct〉c55|850 = 1.821 ± 0.003 for group quenched on tq = 850 °c and 〈cl/ct〉c55|950 = 1.826 ± 0.005 for group quenched on tq = 950 °c. the values have monotonously decreasing trend with increasing tempering temperature tt . chemical composition, tempering and quenching effect general trends were observable on the charts of measured values. the measured l/t ratio values were: • the higher the higher is the concentration of carbon c (see fig. 5), • the higher the lower is the tempering temperature tt (see fig. 4), 49 tomáš zavadil acta polytechnica ctu proceedings steel type c [%] mn [%] si [%] cr [%] ni [%] cu [%] p [%] s [%] c16e/1.1148 > 0.13 > 0.60 > 0.15 < 0.25 < 0.30 < 0.30 < 0.04 < 0.04 (čsn 412020) < 0.20 < 0.90 < 0.40 c35/1.0501 > 0.30 > 0.50 > 0.15 < 0.25 < 0.30 < 0.30 < 0.04 < 0.04 (čsn 412040) < 0.40 < 0.80 < 0.40 c55/1.0535 > 0.52 > 0.50 > 0.15 < 0.25 < 0.30 < 0.30 < 0.04 < 0.04 (čsn 412060) < 0.60 < 0.80 < 0.40 table 1. chemical composition of the steel group used for test samples (composition acc. to czech standards čsn 412020, čsn 412040 and čsn 412060). figure 3. measured l/t ratio values grouped by steel type. c16e (blue), c35 (red) and c55 (green). values for quenching temperatures tq = 850 °c (circle) tend to be in all cases lower than for tq = 950 °c (triangle). • higher for higher quenching temperature tq (see fig. 3). for quenching temperature of tq = 850 °c the values across the chemical composition range shown smaller standard deviation (max s = 0.042%) and across the tempering temperature higher standard deviation (max s = 0.163%). for the quenching temperature tq = 950 °c the values increased to 0.054% for tq = 850 °c and 0.243% for tq = 950 °c respectively. 6. discussion values of the l/t ratio during the experiment on conventional structural steels c16e, c35 and c55 were well controlled within range of 1.812 ÷ 1.835. these steels (especially c16e) were selected for their similarity with to p265gh conventional pressure purpose steel. minimal values were observed in general for low carbon content and low quenching temperatures. the values were changing by heat-induced structural changes and induced phase transformation. that is logical, because they may be a source of a creation and/or change of residual stress. such can be measured by the ultrasonic waves velocity changes or changes of its ratio as explained in chapter 2. a simple relation between the l/t ratio and e.g. the heat treatment temperature however may not be drawn. the reason is that a minimal temperature may be needed to introduce change of stress-strain conditions (i.e. by phase transform etc.). it is expected the l/t ratio as a function of temperature may be only semi-continuous and may include occasional sudden leaps. this phenomena can be demonstrated on samples with higher carbon content and nominal quenching temperature (c55, tq = 950 °c). higher quenching temperature well above ac3 ensures full phase transformation. higher carbon content opens the ttt (timetemperature-transformation) diagram and pushes the bs (bainite start) curve to longer times. original α–fe is transformed to γ–fe and then back with creation of martensitic or martensitic-bainitic structure. low tempering temperature does not break the created martensitic and/or bainitic structure and has minimal impact. because the martensitic phase transformation is accompanied with applied stress from the lattice, the highest l/t ratio values can be found there (〈cl/ct〉c55|950 = 1.833 ± 0.002). other side of the spectra of provided samples represents steel c16e, which has the lowest carbon content and is quenched on temperature tq = 850 °c, i.e. under ac3 curve. it can be expected to observe lower l/t ratio numbers. the phase transformation is only partial (a1 < tq < ac3) and low carbon content pushes the ttt diagram closer to short times, which makes the cooling speed too slow to create purely martensitic structure. additional tempering, especially on higher temperatures, i.e. 600 °c and more tends to create sorbitic structure. such process re50 vol. 19/2018 selection of ultrasonic waves velocities for detection of creep figure 4. measured l/t ratio values grouped by quenching temperatures. c16e (blue) had in general the lowest l/t ratio values, c35 (red) average and c55 (green) the highest for both quenching temperatures tq = 850 °c (circle) and tq = 950 °c (triangle). figure 5. measured l/t ratio values grouped by tempering temperatures. the l/t ratio values decreased for with increasing tempering temperature for all, c16e (max c = 0.2%), c35 (max c = 0.4%) and c55 (max c = 0.6%) and quenching temperatures tq = 850 °c (left) and tq = 950 °c (right). lieves the residual stresses and the l/t ratio is in range of 〈cl/ct〉c16e|850 = 1.816 ± 0.001. all other measurements were in between of those two extremes. the experimental results fitted to the calculated range of cl/ct = 1.73 ÷ 1.87 in chapter 2 for structural steels as was originally expected. the c16e had the most similar chemical composition of the test samples to p265gh conventional pressure purpose steel. this steel is used for production of e.g. membrane walls and first level superheaters. the first approximation of expected value for the p265gh steel unaffected by large-scale heat effect or creep process should be close to the value of this (c16e) steel. expected range acc. to equation (9) and the reference (benchmark) value can be calculated as follows: 〈 cl ct 〉p265|ref = 〈 cl ct 〉j± ± 3 ·maxj √√√√ 1 n − 1 nj∑ i=0 ( cl cti −〈 cl ct 〉j )2 (10) where j is referring to the set of samples of c16e with tq = 850 °c. if the value are filled: 〈 cl ct 〉p265|ref = 1.817±3·0.004 = 1.817±0.012 (11) where 〈cl ct 〉p265|ref is the reference (benchmark) value range for unaffected p265gh steel. notice that the 3s is only 0.7% of the average value. this means the measurement of unaffected steel should be very precise and easily separable from measurements from affected material. the samples subject to severe strengthening shall have values significantly higher than the maximal reference value range. this means the values expected for affected material should fulfill the formula 〈cl ct 〉p265|test � 〈clct 〉 max p265|ref = 1.829. measurements performed on collapsed membrane wall from p265gh pressure purpose steel as described in chapter 3 found maximal values of l/t ratio peaking to 〈cl ct 〉max p265|test = 2.086 close to the crack tip as visible on fig. 2 [13]. that is a value 14% higher than the reference value (it is also 67 times the standard deviation sc16e|850, see formula (10)). it is therefore obvious that severe changes of stress-strain conditions can be clearly separable from the provided reference 51 tomáš zavadil acta polytechnica ctu proceedings value. such changes may be caused by creep, but also other processes, that impact local stress level in the material. 7. conclusion the goal of this article was to find expected benchmark reference value of ratio of longitudinal and transversal ultrasonic waves velocities (l/t ratio) for detection of creep degradation process on conventional non-alloyed pressure purpose steel (e.g. p265gh). the article demonstrates that the l/t ratio is usable method of detection of creep-induced strengthening be measuring induced stress. the performed experiment on three types of structural steel (c16e, c35 and c55) found that the values of l/t ratio are well controlled around the value of 1.82 (dimensionless) for material not subject to degradation and do not drop below 1.812 or raise above 1.835. this is in agreement with computed expected results with range of cl/ct = 1.73÷1.87 for steels with poisson’s ratio ν = 0.25 ÷ 0.30. similarity of c16e structural steel with p265gh pressure purpose steel (steel used for production of membrane walls, first level superheaters etc.) was used to calculate a first level approximation of reference benchmark value range for non-degraded p265gh steel with value 〈cl ct 〉p265|ref = 1.817±0.012. due to the strengthening mechanisms the upper bond of the reference value 〈cl ct 〉max p265|ref = 1.829 is a benchmark for comparison with potentially degraded material. comparison with results measured on collapsed membrane wall from p265gh subject to creep rupture also confirmed the reference value being accurate. it was observed that creep-induced strengthening is both measurable and clearly distinguishable from the reference value, with peak value 〈cl ct 〉max p265|test = 2.086. in conclusion the article proved there is a materialand heat-treatment-specific reference benchmark value that can be used for detection of structural changes of the material through change of its stress-strain conditions. the evaluation of degradation process by this means is expected to be ordinal (comparison with a set of reference samples). such application may be used in case of e.g. on-site rla inspections. continuation of the research will now focus on simplification of the measurement process for on-site inspections and verification of ability to detect the degradation process in its early stages. acknowledgements this article is based on author’s bachelor’s degree thesis and is a part of ongoing postgraduate studies on czech technical university in prague, faculty of nuclear sciences and physical engineering. valuable discussions with mr. jaroslav pitter (materials and corrosion specialist, thesis advisor), colleagues from the atg, ltd., mr. petr žbánek (ut l3), mr. jiří blahušek (ut l3), mr. jan švub (rla inspector) and the professor jiří kunz from the department of materials, faculty of nuclear sciences and physical engineering, ctu in prague, are gratefully acknowledged. this article is a derivative based on a presentations held on the student conference šimáně 2018 and 57th bindt conference in birmingham including respective books of proceedings. references [1] v. sklenička, k. kuchařová, m. svoboda, et al. long-term creep behavior of 9–12%cr power plant steels. materials characterization 51(1):35 – 48, 2003. doi:10.1016/j.matchar.2003.09.012. [2] s. baby, b. n. kowmudi, c. omprakash, et al. creep damage assessment in titanium alloy using a nonlinear ultrasonic technique. scripta materialia 59(8):818 – 821, 2008. doi:10.1016/j.scriptamat.2008.06.028. [3] g. sposito, c. ward, p. cawley, et al. a review of nondestructive techniques for the detection of creep damage in power plant steels. ndt&e international 43(7):555 – 567, 2010. doi:10.1016/j.ndteint.2010.05.012. [4] j. obraz. zkoušení materiálu ultrazvukem. sntl. [5] e. schneider, v. goebbels. vdi-berichte no. 439. [6] s. eldevik, a. a. f. olsen, p. lunde. sound velocity change owing to the acoust-elastic/plastic effect in steel measured using acoustic resonance technology (art). in scandinavian symposium on physical acoustics, geilo, jan 29th feb 1st. [7] m. kobayashi. acoustoelastic theory for plastically deformed solids. jsme international journal ser 1, solid mechanics, strength of materials 33:310–318, 1990. doi:10.1299/jsmea1988.33.3_310. [8] m. kobayashi. ultrasonic nondestructive evaluation of microstructural changes of solid materials under plastic deformation–part i. theory. international journal of plasticity 14(6):511 – 522, 1998. doi:10.1016/s0749-6419(98)00005-9. [9] m. kobayashi. analysis of deformation localization based on proposed theory of ultrasonic wave velocity propagating in plastically deformed solids. international journal of plasticity 26(1):107 – 125, 2010. doi:10.1016/j.ijplas.2009.05.004. [10] a. kumar, t. jayakumar, b. raj, k. ray. correlation between ultrasonic shear wave velocity and poisson’s ratio for isotropic solid materials. acta materialia 51(8):2417 – 2426, 2003. doi:10.1016/s1359-6454(03)00054-5. [11] h. h. hilton, s. yi. the significance of (an)isotropic viscoelastic poisson ratio stress and time dependencies. international journal of solids and structures 35(23):3081 – 3095, 1998. doi:10.1016/s0020-7683(97)00357-0. [12] t. zavadil. ndt assessment of microstructure changes of materials after heat treatment, 2009. bachelor’s thesis, ctu in prague. [13] t. zavadil. the use of ndt methods for detection of microstructure changes of steels when performing residual life assessment of industrial equipment, 2012. master’s thesis, ctu in prague. 52 http://dx.doi.org/10.1016/j.matchar.2003.09.012 http://dx.doi.org/10.1016/j.scriptamat.2008.06.028 http://dx.doi.org/10.1016/j.ndteint.2010.05.012 http://dx.doi.org/10.1299/jsmea1988.33.3_310 http://dx.doi.org/10.1016/s0749-6419(98)00005-9 http://dx.doi.org/10.1016/j.ijplas.2009.05.004 http://dx.doi.org/10.1016/s1359-6454(03)00054-5 http://dx.doi.org/10.1016/s0020-7683(97)00357-0 acta polytechnica ctu proceedings 19:46–52, 2018 1 introduction 2 stress measurement by conventional ultrasonic testing 3 effect of creep on l/t ratio values 4 experimental setup 4.1 samples 4.2 equipment 5 results 6 discussion 7 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.16.0011 acta polytechnica ctu proceedings 16:11–17, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app analysis of the time course of a bridge abutments settlement luboš hruštinec slovak university of technology, faculty of civil engineering, department of geotechnics, radlinského 11, 810 05 bratislava, slovak republic correspondence: lubos.hrustinec@stuba.sk abstract. the paper deals with the time course of the bridge abutments settlement (consolidation of the subsoil) on the motorway d4 in stupava, slovakia. the bridge abutments are founded on an earth embankment 5.5 meters in height and a group of piles. over 6 years of geotechnical monitoring after the construction of bridge abutments, there were measured settlements from 102 to 106 mm. the measured settlement of intermediate bridge piers was only up to 16 mm. geotechnical calculations and analysis are focused on the comparison of the final settlements prognosis and its time course with the real measured values. keywords: bridge abutment, geotechnical calculations, settlement, consolidation. 1. introduction bridge objects no. 211-01 and 211-02 are situated at the crossroads “stupava” of the d4 motorway, which is bridging the d2 motorway in the direction from bratislava to stupava. a three-span bridge structure of total length 100.9 m and width 26.15 m was put into operation in 2010. bridge abutments are founded on piles in an embankment 5.5 m high. on the bridge supports, vertical deformations (settlements) have been continuously measured. in the time period from march 2010 to september 2016 (an overall period of 6.5 years) settlement of the internal piers was measured to be from 12.7 to 16.0 mm and settlement of the bridge abutments from 102 to 106 mm. such an uneven settlement (maximum angular deformation is ∆s / l = 0.0034) causes a significant increase of internal forces in the bearing elements of the bridge structure (especially in the bridge deck) due to the deformation load. in order to avoid failure of the bridge structure, an over-limit uneven settlement had to be eliminated by inserting compensating steel plates into the bridge bearing of the abutments. the paper deals with prognosis of the time course of the subsoil consolidation of the bridge abutments and the causes of their over-limit uneven settlement. this problem is solved in detail in [1]. 2. engineering and geological characteristics of the territory under study engineering and geological conditions of the studied territory and soil properties of the subsoil have been assessed on the basis of archive documentation [2]. the subsoil is formed by quaternary and neogene soils. quaternary soils are represented by deluvial, eolian, proluvial and fluvial sediments. neogene sediments are located in the whole area under quaternary sediments in the form of alternation of the clayey and sandy soils. in the location of bridge structure no. 211 the jm-5 to jm-8 exploration boreholes were conducted to a depth of 15 m. the layers of quaternary fluvial non-cohesive sediments occur in the form of sandy soils. in accordance with stn 72 1001 [3] soils belong to classes: silty sand (s4-sm) and clayey sand (s5-sc). the thickness of the quaternary layer ranges from 0.7 m (borehole jm-5) up to 1.7 m below the surface (borehole jm-6). under the quaternary sandy soils there are neogene clay soils. the clays belong to certain classes: clay with medium plasticity (f6-cl) and clay with high plasticity (f8-ch). the consistency of the clay changes with the increasing depth, from firm to stiff. in the jm-8 borehole, the clay consistency is soft to firm. the groundwater level was found at 0.7 to 1.0 m below the surface. the above mentioned information and knowledge have been taken into account when determining the subsoil model and soil properties used in the geotechnical calculations. 3. determination of the subsoil models and soil properties for bridge abutments no. 1 and no. 4 the subsoil models were defined based on the evaluation of jm-5 and jm-8 boreholes, respectively. computational models of subsoil and bridge abutments used in the geotechnical calculations are shown in figure 1. the depth of the exploration boreholes (jm-5 and jm-8) was 15 m under the original terrain. geotechnical calculations showed that the deformation zone (ha) extends to a depth of 22.3 m below the terrain. based on the evaluation of the results of the geodetic measurements of the settlements and geotechnical calculations, it was assumed in the subsoil models that from the depth of 15 m to the depth of the deformation zone “ha” 11 http://dx.doi.org/10.14311/app.2018.16.0011 http://ojs.cvut.cz/ojs/index.php/app luboš hruštinec acta polytechnica ctu proceedings there are soils of class f8-ch with soft to firm consistency. the soil properties used in the geotechnical calculations are shown in table 1. 4. technical parameters of the bridge structure information about the technical parameters of the bridge structure were taken from [4]. the bridge loadbearing structure (bridge deck) is a monolithic structure made of pre-stressed concrete of class c35/45 with total height of 1.8 m. the three-span bridge beam has a range of fields 26.5 + 36.5 + 26.5 m. the total length of the bridge structure is 100.9 m (figure 2). in the transverse direction, the bridge consists of two separate structures. for the right traffic lane, the bridge is marked as no. 211-01 and for the left traffic lane no. 211-02 (figure 3). in the cross-section, two pre-stressed concrete beams are connected with a reinforced concrete slab (bridge deck). the total width in the transverse direction of the bridge is 13.75 m (structure no. 211-01), and 11 m (structure no. 21102). geometric boundary conditions of the abutment structures no. 1 and 4 (including embankment and backfill) for structures no. 211-01 and no. 211-02 have been taken from the project documentation[4]. bridge abutments are deep founded on piles and on the embankment (figure 2). the foundation consists of a group of piles with a diameter of 0.9 m, a length of 14 m (no. 211-01 group of 7 piles and no. 211-02 group of 6 piles). computational models of bridge abutments no. 1 and 4 are shown in figure 1. 5. determination of load states for geotechnical calculations the loads due to the bridge structures were taken from the static analysis mentioned in [1]. the total loads acting on the bridge abutments and subsoil are shown in table 2. in the geotechnical calculations, the most significant loads of the subsoil from the individual structures (embankment, abutment and the backfill behind the abutment) and the technological progress of the construction of bridge structures no. 211-01 and 211-02, were considered. three following load states were defined: • 1st load state (1st ls): construction of the embankment below the abutments. under the abutment no. 1 the height of the embankment is hn1 = 5.5 m, which causes a uniform plane load qn1 = γ × hn1 = 20 × 5.5 = 110 knm−2. under the abutment no. 4 the height of the embankment is hn4 = 4.8 m, which causes a uniform plane load qn4 = γ × hn4 = 20 × 4.8 = 96 knm−2. • 2nd load state (2nd ls): construction of piles, abutments and bridge decks (table 2). • 3rd load state (3rd ls): construction of a backfill behind the bridge abutments. behind the abutments the height of the backfill is hz = 4.3 m, which causes a load of size qz = γ × hz = 20 × 4.3 = 86 knm−2. from the defined load states it follows that the total load of the subsoil at the level of the terrain from embankments and backfills, causes the load of up to q1 = 196 knm−2 (for abutment no. 1) and q4 = 182 knm−2 (for abutment no. 4). the total vertical load acting locally below the abutments (at the foundation gap) has a size q211−01 = 220.9 knm−2 (for abutment no. 211-01) and q211−01 = 232.5 knm−2 (for abutment no. 211-01). from these values of load intensity it is obvious that the large embankment and backfill will have a significant impact on the final settlement of the abutments and cannot be neglected in the calculations. 6. results from geodetical measurements of the bridge supports settlement the vertical displacements (settlements) of the bridge supports have been continuously measured using the geodetic method of a very precise levelling. in the period from july 13, 2010 to september 3, 2016 (i.e. over 6 years), a total of 23 stage measurements have been performed. the detailed results of the geodetic measurements are given in [1]. graphical evaluation of the time course of settlement of supports no. 1 to 4 of the bridge structure no. 211-01 and no. 211-02 is shown in figure 4. the measured maximum settlements of the middle piers were 12.7 to 16.0 mm and the settlements of the edge bridge abutments were 102.0 to 106.0 mm. from the static analysis of a bridge structure, the maximum value of the settlement difference of the neighbouring supports ∆s = 20 mm [1] resulted. the measured maximum settlement difference of the neighbouring supports is ∆s = 106– 16 = 90 mm and maximal uneven settlement is ∆/l = 90/26500 = 0.0034. the above-mentioned overlimit uneven settlement causes a significant increase of the internal forces in the bridge bearing elements (especially in the bridge deck) due to the deformation load. to compensate for the over-limit uneven settlements of the bridge supports, the bridge deck structures at the place of abutments were lifted and steel washers with a thickness of 45 mm were inserted on november 15, 2010. 7. evaluation of geotechnical calculations the geotechnical calculations were focused on the determination of the rate of the final settlement and the time course of settlements of the bridge abutments no. 1 and 4. the calculations were performed in accordance with valid standards [5, 6]. the final settlement was calculated using the finite element method (planar task) and the time course of settlements was analyzed in terms of the theoretical assumptions of 12 vol. 16/2018 bridge abutments settlement geological period soil class-symbol consistency γ [kn/m3] ν [-] e [mpa] cv [m2·s−1] quaternary silt with low plasticity f5-ml firm 20.0 0.40 4.0 5.0e-06 silt sand s4-sm fine fraction is firm 18.0 0.30 8.0 7.5e-05 clayey sand s5-sc fine fraction issoft firm 18.5 0.35 8.0 3.8e-06 neogene clay with medium plasticity f6-ci soft to firm 21.0 0.40 3.0 5.4e-07 clay with high plasticity f8-ch soft to firm 20.5 0.42 2.0 2.2e-07 clay with high plasticity f8-ch firm 20.5 0.42 3.0 2.2e-07 clay with high plasticity f8-ch firm to stiff 20.5 0.42 4.0 2.2e-07 clay with high plasticity f8-ch stiff 20.5 0.42 6.0 2.2e-07 table 1. soil properties used in geotechnical calculations. figure 1. computational models for abutments no. 1 and 4 used in geotechnical calculations. 13 luboš hruštinec acta polytechnica ctu proceedings figure 2. longitudinal cross section of the bridge structure [4]. figure 3. cross section of the bridge structure [4]. bridge structure abutment type of load vz[kn] hx [kn] hz [kn] mx [knm] my [knm] 211-01 1 and 4 dead weight 7,870.5 0.0 0.0 0.0 2,245.0 211-02 1 and 4 dead weight 6,685.3 0.0 0.0 0.0 1,952.3 table 2. loads for the bridge abutments used in geotechnical calculations [4]. figure 4. the measured time course of settlements (si) of the bridge supports (piers and abutments) and settlements difference (∆si) of neighbouring supports. 14 vol. 16/2018 bridge abutments settlement figure 5. calculated time course of settlement of the bridge structure no. 211-01, abutment no. 1. bridge structure abutment calculated final settlement s measured settlement (3. 9. 2016) settlements (difference calc.-meas.)abutment embankment backfill completely (embankment, abutment, backfill) sa [mm] se [mm] sb [mm] sc [mm] sm [mm] ∆sc [mm] 211-01 1 26.65 92.00 35.45 154.10 102.00 52.10 4 34.45 86.60 40.70 161.75 106.00 55.75 211-02 1 28.30 93.15 36.20 157.65 104.70 52.95 4 32.90 85.30 38.25 156.45 104.00 52.45 table 3. comparison of calculated and measured settlements with the prognosis of further settlement ∆sc [mm]. one-dimensional consolidation [7]. geotechnical calculations yielded many quantitative and qualitative results. the detailed results of the calculations can be found in the author’s archive records. this article presents only some representative results of geotechnical calculations and their processing in a clear tabular and graphical form. graphical evaluation of the calculated time course of settlement of the individual structural parts (embankment, abutment, backfill, and the complete structure) of the abutment no. 1 (bridge no. 211-01) is shown in figure 5. the table comparison of the measured (as of september 3, 2016) and the calculated values of the final settlements of the bridge abutments with the prognosis of expected increase of settlement ∆s is given in table 3. the relative percentage comparison with the calculated final settlement is presented in table 4. the graphical evaluation of the comparison of the measured settlements and the calculated final settlements of bridge structure no. 211-01 for a period of 0 to 10 years is shown in figure 6. from the evaluation and analysis of the results of the geotechnical calculations and a comparison with the measured settlements, the following facts and knowledge follow: • calculated values of the final settlements (average values in the centre of abutments) range from 154.1 mm (bridge no. 211-01, abutment no. 1) to 161.75 mm (bridge no. 211-01, abutment no. 4). from the total value of the final settlements, the effects of loads due to the individual structural parts is the following: due to the embankment up to 59.7 % (from 54.5 to 59.7 %), due to the backfill behind the abutment it is up to 25.2 % (from 23.0 to 25.2 %) due to the subsoil loading by the abutment structure it is to 21.3 % (from 17.3 to 21.3 %). from the above mentioned it follows that the final value of the settlement is mainly affected by loading of the subsoil by the embankment and the backfill behind abutment. the final settlement caused by the embankment and backfill causes up to 82.7 % (from 78.7 to 82.7 %) of the total final settlements of the bridge abutments. the size of the calculated settlements is also significantly affected by the presence of clays soils with soft consistency in the bridge abutments subsoil. • the calculations results of the time course of settlements (under the assumed boundary conditions) showed that the subsoil consolidation will be finished (consolidation degree u = 100 %) after 51.1 years (figure 5), or 80 % degree of consolidation will be achieved after 10 years (figure 6). • from the comparison of the measured settlements (as of september 3, 2016) and the calculated final settlements, there is the result that the measured values reach up to 66.5 % of the final settlements, i.e. approximately 2/3 of the final settlements value (table 4). the expected next settlements ∆s of the bridge abutments are in the range from 52.1 mm to 15 luboš hruštinec acta polytechnica ctu proceedings bridge structure abutment calculated final settlement s measured settlement (3. 9. 2016) settlements difference (calc.-meas.)abutment embankment backfill completely (embankment abutment, backfill) so [%] sn [%] sz [%] sc [%] sm [%] ∆sc [%] 211-01 1 17.3 59.7 23.0 100.0 66.2 33.8 4 21.3 53.5 25.2 100.0 65.5 34.5 211-02 1 18.0 59.1 23.0 100.0 66.4 33.6 4 21.0 54.5 24.0 100.0 66.5 33.5 table 4. the percentage comparison of calculated and measured settlements. figure 6. time course of settlement of the bridge structure no. 211-01, abutments no. 1 and 4 – comparison of measured and calculated values over a 10-year period. 55.75 mm (table 3), i.e. 33.5 to 34.5 % of the final settlements (table 4). the above findings must be taken into account in the complex assessment of the reliability of the bridge structures no. 211-01 and no. 211-02. 8. conclusions based on the analysis and evaluation of the geotechnical calculations undertaken it can be stated that the over-limit settlements of the bridge abutments occurred mainly due to the following factors: • failures during the execution of the geotechnical survey (mainly insufficient depth of survey boreholes that did not reach the depth of the deformation zone; ending boreholes in cohesive soils of soft consistency; insufficient determination of geotechnical data valid for soils in the subsoil). • failures during the execution of the geotechnical calculations and design of the bridges abutments (especially neglecting the significant part of the loading of the subsoil due to the embankment and the backfill bridge abutments; neglecting the consolidation processes in the subsoil). to ensure the reliable operation of bridge objects no. 211-01 and no. 211-02, the following recommendations and proposed measures were formulated: • at regular intervals, to visually check the bridge structures and their load-bearing elements. during the bridge checking it is necessary to focus on deformation of the structures (including the embankment and the surrounding terrain), the origins of cracks and other disproportional phenomena (deformation of the terrain, inadequate deformations and rotation, etc.). • to perform ongoing geodetic measurements of vertical displacements (settlements) of bridge supports no. 1 to 4 (including internal piers no. 2 and no. 3) at a minimum of twice per year. • after each periodic measurement it is necessary to graphically evaluate the measured values of settlement and to compare them with the calculated values. if the measured settlement values are not 16 vol. 16/2018 bridge abutments settlement in accordance with the expected values, it is necessary to take measures to eliminate the uneven settlements and their effects on the reliability of the bridge structure. this involves following the recommendations of the static analysis of the bridge structure, such as e.g., inserting washers into the abutments bridge bearings. • if the measured settlements are higher than the expected ones, or they progressively rise, it is necessary to carry out an additional geotechnical survey to the depth of the deformation zone. • in case of another disproportional progressive increase of the settlements, it is necessary to take more complex measures, such as e.g., pre-injection of the subsoil using jet grouting technology. list of symbols γ unit weight with natural moisture [kn m−3] ν poissin ratio [–] edef modulus of deformation [mpa] cv coefficient of consolidation [m2 s−1] v y vertical force [kn] hx horizontal force in the x-axis direction [kn] hy horizontal force in the y-axis direction [kn] mx moment in the x-axis direction [knm] mx moment in the y-axis direction [knm] s settlement; vertical displacement [mm] ∆s settlement difference [mm] u degree of consolidation [%] acknowledgements the author is grateful for support from the grant agency vega of the slovak republic, project no. 1/0412/18. references [1] j. halovonik. assessment of stress in the bridges structures no. 211-01 and 211-02 – crossroad stupava. bratislava, 2016. [2] i. modlidba. crossroad stupava – south on the d2 motorway. a detailed engineering-geological survey. final report. terratest. bratislava, 2006. [3] stn en 72 1001: classification of soil and rock in engineering geology and geotechnics, 2010. [4] m. sloboda. crossroad stupava – south on the d2 motorway. documentation of actual construction. strabag – doprastav, 2011. [5] stn en 1997-1 eurocode 7: geotechnical design. part 1: general rules. 2005. [6] stn 73 1001: geotechnical structures. foundation, 2010. [7] j. jesenák. soil mechanics. stu bratrislava, 1994. 17 acta polytechnica ctu proceedings 16:11–17, 2018 1 introduction 2 engineering and geological characteristics of the territory under study 3 determination of the subsoil models and soil properties 4 technical parameters of the bridge structure 5 determination of load states for geotechnical calculations 6 results from geodetical measurements of the bridge supports settlement 7 evaluation of geotechnical calculations 8 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.20.0056 acta polytechnica ctu proceedings 20:56–64, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app thermo-mechanical fatigue analysis of a steam turbine shaft martin nesládeka, ∗, josef jurenkaa, michal bartošáka, milan růžičkaa, maxim lutovinova, jan papugaa, radek procházkab, jan džuganb, petr měšťánekc a czech technical university in prague, faculty of mechanical engineering, technická 4, 166 07 prague 6, czech republic b comtes fht a.s., průmyslová 995, 334 41 dobřany, czech republic c doosan škoda power s.r.o., tylova 1/57, 301 28 plzeň, czech republic ∗ corresponding author: martin.nesladek@fs.cvut.cz abstract. increasing demands on the flexibility of steam turbines due to the use of renewable energy sources substantially alters the fatigue strength requirements of components of these devices. rapid start-ups as well as the increased number of the load cycles applied to the turbines must be handled by design methodologies. the goal of the work presented in this paper was to provide a computational framework applicable to the thermo-mechanical fatigue (tmf) prediction of steam turbine shafts. the so-called damage operator approach by nagode et al. has been implemented to the software codes and applied to fatigue analysis of the thermo-mechanical material response computed numerically by the finite element analysis. experimental program conducted in order to identify the material thermo-mechanical behavior and to verify numerical simulations is introduced in the paper. some results of tmf prediction of a sample steam turbine shaft are shown. keywords: fatigue prediction, thermo-mechanical fatigue, finite element analysis, chaboche elasticplastic model. 1. introduction the energy industry has been experiencing significant changes recently. the overall share of renewable sources in the european union countries has reached 17 percent [1]. the leading country in this statistic is sweden having more than 50 percent of energy from renewable sources. in czech republic it is approximately 15 percent and the overall trend is that this share will continue to grow significantly in the future the eu strategy is to reach at least 27% by 2030. however, instability of the power output due to the nature of the renewable sources may cause problems in energy supplies if it is not compensated by some back-up sources. currently, there are no reliable ways to save energy for power shortages in the network in a sufficiently large scale. this role must be taken over by traditional thermal power sources such as gas-fired power plants providing "just-in-time" energy supplies. the problem is that these plants already have to or will have to cover shortages in the daily regime. increased number of start-ups and shut-downs as long as increasing demands on speed-up of the start-ups set new challenges in the gas and steam turbine design. components of steam turbines are often exposed to thermo-mechanical loading due to variable thermal and mechanical loading conditions induced by the turbine start-ups and shut-downs. the main damage mechanisms, which if they occur together constitute the therm "thermo-mechanical faigue" (tmf), are often referred to as oxidation, mechanical fatigue and creep [2, 3]. constitutive material model derived from the material testing, finite element analysis and suitable fatigue criterion is required for the reliable damage assessment of components subjected to cyclic thermo-mechanical loading conditions. this work is focused on the tmf assessment of a sample case of a steam turbine shaft. motivation for this work originates from the need for revising the current steam turbine design strategies due to the above-mentioned reasons. classical low-cycle fatigue (lcf) criteria as smithwatson-topper (swt) [4] and the nihei [5] methods in their original form can only cope with isothermal loading. for the case of thermo-mechanical loading, cycle reference temperature, which could be used to select fatigue curves for the isothermal models, is somewhat arbitrary and error-prone. the so-called damage operator approach (doa) by nagode et al. [6, 7] overcomes this issue by introducing a procedure for continuous damage parameter calculation. in the original form it is based on the swt method, however the damage induced by nonstationary thermal and mechanical loading is evaluated in a chronological manner, i.e. damage is computed and accumulated in any time point no matter if it is load reversal or not. such approach allows for handling arbitrary thermal and mechanical load 56 http://dx.doi.org/10.14311/app.2018.20.0056 http://ojs.cvut.cz/ojs/index.php/app vol. 20/2018 tmf analysis of a steam turbine shaft sequences without the need for evaluating the abovementioned cycle reference temperature. the work presented here was done within the flexturbine project, which aims to revise the design of key turbine components and propose new design methodologies so that the turbines can cope with the changing situation in the transmission grids. the paper describes the approach to prediction of thermomechanical fatigue of steam turbine shafts. especially the high-pressure sections of turbine shafts are exposed to complex thermo-mechanical loading due to variable service conditions. within the experimental part of the work, experiments for identifying parameters of the selected constitutive model as well as parameters of damage model were carried out. the assumed constitutive model of the material mechanical response is temperaturedependent elastic-plastic and for the purpose of the structural finite element analysis (fea) parameters of the chaboche non-linear kinematic hardening were identified. the doa allows for processing stress histories computed by fea in order to perform correction on the material viscous behavior, if needed. strain controlled fatigue tests with hold period may be used for finding the parameters of the maxwell model of viscoplasticity, which may then be applied for evaluating the viscous strain and subsequent stress relaxation. a set of tmf tests for verifying the prediction methodology have been scheduled. various phasing of temperature and mechanical load is tested. the experiments are conceived as strain-controlled with two channels of strain applied to tubular specimens axial and torsional. tmf test setup is briefly introduced in the paper and some preliminary results are presented. thermo-mechanical stress-strain response from the tests and numerical simulations is compared. the elastic-plastic material model is used in the structural fea of the steam turbine shaft, which is performed on the basis of the simulated temperature fields that are obtained from the previous transient thermal fea. finite element calculations for the steam turbine shaft are performed in the commercial finite element software. subsequently, tmf is predicted for the turbine shaft on the basis of the calculated stress, strain and temperature histories. 2. experimental work as mentioned within the introduction section, the experimental work may be divided into two main groups depending on the general purpose of the individual test on the one hand basic material tests to identify parameters of constitutive models and on the other hand tests intended for verifying the proposed modeling methodology. the material subjected to testing was a chrome-molybdenum creep resistant steel applicable to steam turbine shafts. figure 1. the temperature dependent young’s modulus. figure 2. the temperature dependent cyclically stable cyclic stress-strain curves. figure 3. the temperature dependent manson-coffin and basquin curves. 57 m. nesládek, j. jurenka, m. bartošák et al. acta polytechnica ctu proceedings (a) . detail of grips with the test specimen and instrumentation (b) . overall view of the tmf test setup (c) . tubular test specimen with surface speckle pattern and welded thermocouple figure 4. tmf test setup and specimen. 1 servohydraulic axial-torsional test machine, 2 resistance heating, 3 camera of the dic system, 4 infrared camera, 5 test specimen. 2.1. basic mechanical material tests static tensile tests as well as lcf tests under the isothermal conditions have been carried out. the assumed temperature range of the tests was 20 – 600 ◦c, which correspond to operating conditions of the sample steam turbine. circular specimens with 10 mm diameter in the gauge section have been tested. radiant heating of specimens was controlled by a split furnace. as a minimum, ten samples have been tested on each curve. fig. 1, 2 and 3 show the main results of the isothermal material tests, i.e. temperature dependent normalized young’s modulus, cyclic stress-strain curves (csscs) and the manson-coffin and basquin curves, respectively. the young’s modulus and the csscs have been used for identifying the elastic-plastic material model based on the chaboche’s procedure section 3. the temperature dependent manson-coffin and basquin curves are used as an input to the doa. numerical procedure based on this approach is described in section 4. lcf strain-controlled tests with hold period have been performed in order to identify temperature de0 500 1000 1500 0.2 0.4 0.6 0.8 1 1.2 t [s] n o rm a li z e d s tr e s s [ − ] 450 ° c 550 ° c 600 ° c figure 5. the temperature dependent stress relaxation curves. pendent stress relaxation curves fig. 5. this is an important input to the doa allowing for separating elastic-plastic and viscous strain and modeling stress relaxation section 4. note that the curves in fig. 5 are shown for illustrative purposes and will be further refined with continuation of the experimental program. 58 vol. 20/2018 tmf analysis of a steam turbine shaft figure 6. time dependence of axial strain and temperature during the in-phase tmf test. furthermore, creep tests should be conducted in order to identify constants of the larson-miller parameter that is applied to evaluating creep damage. 2.2. verification tests experimental setup for tmf tests is shown in fig. 4a and 4b. the test machine is equipped with a servohydraulic actuator allowing for combined axial and torsional loading. resistance heating is applied to specimens and the induced temperature is controlled by a thermocouple welded in the center of the specimen gauge section. online axial and torsional deformation control is done by an optical measurement system based on the digital image correlation (dic). tubular specimens with 8.5 mm in diameter and 0.75 mm wall thickness of the gauge section are sprayed with temperature-stable paint creating a random speckle pattern fig. 4c. design of the test specimens conforms to astm e2207 standard. triangular strain and temperature in-phase and out-of-phase loading is assumed within the tests fig. 6 and 7. until now, tests with axial strain loading at εa = 0.007 and r = −1 have been conducted. the maximum test temperature is 600 ◦c. more tests combining axial and torsional loading with various phases are scheduled in order to check the prediction quality of the selected model. 3. numerical simulations 3.1. constitutive material model a constitutive material model must be able to describe the cyclic material behavior for the service temperature range of the steam turbine shaft. the temperature-dependent chaboche non-linear kinematic hardening model [8, 9] is selected for the elasticplastic fea. the yield criterion may be expressed as f(σ−α) = k, (1) where k is the material temperature dependent yield stress, σ is the stress tensor and α denotes the backfigure 7. time dependence of axial strain and temperature during the out-of-phase tmf test. stress tensor. the function f() denotes the von mises stress norm, which may be in this case evaluated as f(σ−α) = √ 3 2 (s −α′) : (s −α′), (2) where s and α′ are the deviatoric parts of the stress and the backstress tensors, respectively. the : operator denotes the tensor contraction. assuming the non-linear kinematic hardening model, the total backstress is composed of multiple backstress components as follows: α = ∑ i αi (3) the non-isothermal evolution of the backstress components is defined as: α̇i = 2 3 ciε̇p −γiαiλ + 1 ci ∂ci ∂t αiṫ. (4) ci and γi are temperature dependent material parameters, which may be derived from the material mechanical cyclic stress-strain response. εp is the plastic strain tensor and λ is accumulated plastic strain. in our case, the model with three backstress terms was selected, where γ3 is assumed constant and equal to zero. plots in fig. 8 and 9 show comparison of the measured temperature-dependent material mechanical response with the one simulated by finite elements using the above-mentioned constitutive model. the plots show good agreement of simulated and experimentally determined material response. 3.2. fea of a steam turbine shaft the non-linear transient elastic-plastic fea of a sample steam turbine shaft has been carried out in order to investigate local thermo-mechanical material response due to the turbine load regimes. in this paper, results of the cold-start load scenario are presented. at first, transient thermal analysis has to be conducted in order to analyze the time-dependent temperature spatial 59 m. nesládek, j. jurenka, m. bartošák et al. acta polytechnica ctu proceedings figure 8. comparison of measured and simulated mechanical response due to in-phase uniaxial mechanical and thermal loading. figure 9. comparison of measured and simulated mechanical response due to out-of-phase uniaxial mechanical and thermal loading. distribution. the cold-start regime is characterized by the turbine start-up from ambient temperature and after shut-down it cools down back to almost 20 ◦c. so the service temperature range corresponds to ∼ 20 − 600 ◦c and especially during the start-up process the material exhibits severe temperature gradients in the subsurface layers since it is exposed to superheated steam. the heating rates are the matter of the individual turbine design that has to ensure that the material fatigue strength corresponding to the expected turbine life is not exceeded. three basic turbine load regimes are assumed in the design cold, warm and hot starts differing, from a material point of view, in the initial temperature and allowable heating rates. however, regarding the necessary computation strategy, these regimes are similar. temperature distribution in the turbine is a function of complex time evolution of heat flux governed by the material surface and steam flow difference and heat transfer coefficients. the induced thermal gradients result in the constrained material expansion with pronounced thermal stresses. furthermore, mechanical loads due to shaft rotation and steam flow pressure on blade airfoils have to be taken into account. fig. 10 shows time and subsurface distribution of thermo-mechanical material response within a single cold-start cycle in the domain marked in fig. 10a. the observable discontinuities in temperature are present due to the step changes in the boundary conditions, which partly reflect the nature of the service conditions. on the other hand, this also partly results from modeling imperfections caused by higher analysis incrementation than would be suitable to cover some steep transients. however, this is a reasonable compromise between accuracy and solution cost. the plots in fig. 10 illustrate how complex the material thermo-mechanical response could be in such structure. for instance, in the r3 notch root the cyclic plastic stress-strain state under variable temperature is present. this justifies selection of the doa as a method to assessing the tmf damage. 4. thermo-mechanical fatigue prediction fig. 11 graphically presents the main steps, inputs and outputs of the doa, which was adopted and implemented to software codes for the purpose of this work. as mentioned in the introduction section, the main advantage of this approach is that it treats thermo-mechanical material response in step-by-step manner (i.e. in increment-by-increment manner if translated into the finite element terminology). this allows for handling arbitrary stress-strain and temperature histories in tmf prediction, while assigning the proper parameters of the material damage model to the individual load time points. in general, it requires local elastic-plastic stress and temperature from fea as an input. the material inputs are temperature-dependent csscs, mansoncoffin and basquin curves, larson-miller parameter and stress relaxation curves for cases, where the material viscous effect is substantial. note that d = 1/n denotes mechanical fatigue damage per single cycle. prandtl hysteresis operator is employed in order to model the continuous fatigue mechanical damage influenced by temperature changes. the total damage is evaluated as summation of the mechanical part and the creep part computed separately. the oxidation effect is treated indirectly since the material fatigue and creep tests are performed under environmental conditions. palmgren-miner’s rule [10] and robnson’s life-fraction rule [11] are used for accumulating the mechanical fatigue and creep damage, respectively. this procedure has been coded to an in-house c++ procedure allowing for processing thermo-mechanical response from fea. refer to [12] for more details. 5. results and discussion tmf prediction results obtained by the doa are shown in fig. 12 and tab. 1. the contour plot in fig. 12b shows distribution of 6 √ d quantity, where d denotes the tmf damage due to the cold-start turbine load regime. it may be observed that the critical place is the r3 domain. note that these results were obtained by the doa without taking the 60 vol. 20/2018 tmf analysis of a steam turbine shaft (a) . the r3 domain and direction of the nodal path from which the results below have been extracted. (b) . time and spatial distribution of thermo-mechanical material response in the r3 domain. figure 10. fea results in the r3 domain of the sample steam turbine shaft. r2 r3 r4 r5 hp1 upper fillet hp1 lower fillet tmf damage 1.31 · 10−4 2.66 · 10−4 1.69 · 10−4 2.11 · 10−4 2.25 · 10−5 1.81 · 10−4 tmf life 7633 3759 5917 4739 44 444 5525 table 1. tmf fatigue prediction results by the doa in the selected domains of the steam turbine. 61 m. nesládek, j. jurenka, m. bartošák et al. acta polytechnica ctu proceedings strain controlled nonlinear maxwell model to compute the viscoplastic strain εvp thermo-mechanical elastic-plastic fea by ansys mechanical strain response modelling by using the prandtl hysteresis operators. assumes kinematic strain hardening. t1 t2 t3 cyclic stress-strain curves σa εa tinterp ˙εvp = ( σ − k(t) k(t) )n(t ) t1 t2 t3 stress relaxation curves σ t tinterp elastic-plastic strain calculation εep(t) = ε(t) − εvp(t) for t(t) exceeding the material creep temperature tc ε(t) relaxed stress calculation using the prandtl hysteresis operators. εep(t) t1 t2 t3 cyclic stress-strain curves σa εa tinterp t1 t2 t3 tinterp low-cycle fatigue curvespsw t d = 1/n fatigue damage estimation (low-cycle fatigue damage operator approach) df creep damage estimation by using the larson-miller parameter dc σ(t), t(t) εep(t) σ(t) df dc d = dc + df total damage creep master curve σ lmp 1. 2. 3. 4. 5. 6. 7. material inputs doa processes inputs from fea figure 11. flowchart to illustrate the basics of the damage operator approach by nagode et al. 62 vol. 20/2018 tmf analysis of a steam turbine shaft (a) . steam turbine domain nomenclature adopted in this paper. (b) . sample contour plot of tmf damage distribution analyzed by the doa in the main sealing part of the steam turbine. figure 12. tmf prediction by the doa. material viscoplasticity into account. the difference between these values and those presented in [12] is due to different approach to handling the stress-strain histories. the results presented in this paper have been evaluated by using only the signed von mises stress from the fea as an input to the doa. the mechanical strain values have been obtained by using the prandtl hysteresis operators, which are part of the doa fig. 11. this approach gives a stress-strain response that more closely matches the physical behavior of the material than the approach that had been used in [12]. in that previous paper, the signing was applied to both the stress and the strain histories extracted from fea. this naturally may lead to distorted hysteresis loops and may have resulted to, for instance, different stress means and amplitudes. if these stress and strain histories are used as a direct input to the function evaluating the tmf damage, different results may be obtained as well. 6. conclusions this paper briefly presents the experimental program conducted in order to provide the necessary material and verification data for the adopted tmf approach. the tmf experimental setup was introduced and the experimentally determined stress-strain response was compared to the simulated one. the comparison shows good agreement of both experimental and numerical data. the doa in the form allowing for the viscous correction of strain histories and stress relaxation has been implemented to the software codes and the results of the tmf prediction on the sample steam turbine shaft are presented in the text. acknowledgements this project has received funding from the european union’s horizon 2020 research and innovation program under the grant agreement no. 653941. references [1] share of renewables in energy consumption in the eu reached 17% in 2016. http://ec.europa.eu/eurostat/ documents/2995521/8612324/8-25012018-ap-en.pdf/ 9d28caef-1961-4dd1-a901-af18f121fb2d. accessed: 2018-07-30. [2] r. w. neu, h. sehitoglu. thermomechanical fatigue, oxidation and creep: part i. damage mechanisms. metall trans a 20(9a):1755–1767, 1989. [3] r. w. neu, h. sehitoglu. thermomechanical fatigue, oxidation and creep: part ii. life prediction. metall trans a 20(9a):1769–1783, 1989. [4] k. n. smith, p. watson, t. h. topper. a stress-strain function for the fatigue of metals. j mater, jmlsa 5(4):767–778, 1970. [5] m. nihei, p. heuler, c. boller, t. seeger. evaluation of mean stress effect on fatigue life by use of damage parameters. int j fatigue 8(3):119–126, 1986. 63 http://ec.europa.eu/eurostat/documents/2995521/8612324/8-25012018-ap-en.pdf/9d28caef-1961-4dd1-a901-af18f121fb2d http://ec.europa.eu/eurostat/documents/2995521/8612324/8-25012018-ap-en.pdf/9d28caef-1961-4dd1-a901-af18f121fb2d http://ec.europa.eu/eurostat/documents/2995521/8612324/8-25012018-ap-en.pdf/9d28caef-1961-4dd1-a901-af18f121fb2d m. nesládek, j. jurenka, m. bartošák et al. acta polytechnica ctu proceedings [6] m. nagode, f. l angler, m. hack. a time-dependent damage operator approach to thermo-mechanical fatigue of ni-resist d-5s. int j fatigue 33(5):692–699, 2011. doi:https://doi.org/10.1016/j.ijfatigue.2010.11.009. [7] m. nagode. continuous damage parameter calculation under thermomechanical random loading. methodsx 1:81–89, 2014. https://doi.org/10.1016/j.mex.2014.07.004. [8] j. l. chaboche. time-independent constitutive theories for cyclic plasticity. int j plasticity 2(2):149–188, 1986. [9] j. l. chaboche. constitutive equations for cyclic plasticity and cyclic viscoplasticity. int j plasticity 5(3):247–302, 1989. [10] a. palmgren. die lebensdauer von kugellagern. zeitschrift des vereines deutscher ingenieure 68(14):339–341, 1924. [11] e. l. robinson. effect of temperature variation on the long-time rupture strength of steels. trans asme 74:777–780, 1952. [12] m. nesládek, m. bartošák, j. jurenka, et al. thermo-mechanical fatigue prediction of a steam turbine shaft. matec web conf 165, 2018. https://doi.org/10.1051/matecconf/201816522016. 64 http://dx.doi.org/https://doi.org/10.1016/j.ijfatigue.2010.11.009 https://doi.org/10.1016/j.mex.2014.07.004 https://doi.org/10.1051/matecconf/201816522016 acta polytechnica ctu proceedings 20:56–64, 2018 1 introduction 2 experimental work 2.1 basic mechanical material tests 2.2 verification tests 3 numerical simulations 3.1 constitutive material model 3.2 fea of a steam turbine shaft 4 thermo-mechanical fatigue prediction 5 results and discussion 6 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2017.13.0152 acta polytechnica ctu proceedings 13:152–156, 2017 © czech technical university in prague, 2017 available online at http://ojs.cvut.cz/ojs/index.php/app the occurrence of synchronized jumping crowd at sports stadiums martin verner czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic correspondence: martin.verner@fsv.cvut.cz abstract. since the age of the rome empire, people have been going for fun at the stadiums. today, important reasons for visiting the stadium are sport games or cultural performances. the text is about occurrence of specific types of behaviour during sport games. the experiment was realized at four different sports stadiums and during three kinds of sports. the worst case of loads for the grandstand is the synchronized jumping of a crowd. this is type of behaviour, which can cause over limit vibration, visitors discomfort on the tribune, in the extreme case collapse of the grandstand. in any country today, the standards or references or math models are not for application of dynamic load to grandstand. only the british standards are making recommendation how the designer should work with dynamic loads. keywords: grandstands, synchronized jumping, football, ice-hockey, biathlon. 1. introduction the jumping (especially the synchronized jumping of all rows of fans synchronized) is the main element of fans supporting their lovely team at stadiums. in 2010, the grandstands were shaking at football stadium in the nuremberg, germany. behaviour of spectators caused “huge” vibration of grandstands during match of bundesliga between 1. fc nuremberg and fc eintracht frankfurt. the tribune survived the rising level of vibrations [1]. the main question of the text is: what is the frequency of the occurrence of synchronized jumping crowd at sports stadiums in the czech republic. the research took place in the years 2013–2017. there was monitored the fans behaviour during the different sports matches and the acceleration of the grandstand structure was measured. the experiments were realized at the football stadium – prague letná, ice-hockey stadium – pilsen, football stadium – prague eden and biathlon stadium – new town in moravia. example types of construction figure 3, 4. 2. types of fans behaviour during the experiment, nine types of fans behaviour was recorded: walking and running, jumping, bouncing/bobbing/jouncing, swaying, handclapping, static state, goal scoring and mexican wave. the response to the grandstand vibrations are the higest during jumping. the type of behaviour -jumpingis divided to four subtypes [2]. jumping of the 1st type: the fans are jumping alone independently. the synchronization is realized only by drum. the drum is usually placed between fans named "ultras". figure 1. photo of fans: football stadium – prague letna(upper) [2], football stadium – prague, eden (lower) [3]. 152 http://dx.doi.org/10.14311/app.2017.13.0152 http://ojs.cvut.cz/ojs/index.php/app vol. 13/2017 the occurrence of synchronized jumping crowd figure 2. photo of fans: ice-hockey stadium – pilsen (upper) [4], biathlon stadium – new town in moravia (lower) [5]. jumping of the 2nd and the 3rd types: the fans are grabbing around shoulders and they are jumping all rows (vertically (type 2) or horizontally (type 3)). the synchronization is the fastest of all types of behaviours and vibrations on the grandstand are maximal. jumping of the 4th type: the type is called: sit and jump. spectators are sitting then they jump up when the ultras leader order to do it. the text focuses on the 2. and the 3. types of jumping. these types of fans behaviours caused the maximal vibration of the grandstand. 3. realized experiments 3.1. experiment new town in moravia the fans of behaviour (figure 2) influenced by the instability of the assembled grandstands in the new town in moravia. the jumping was prohibited because the organizers were afraid of high vibrations of the grandstands during this type of behaviour. interdiction of the jumping was respected during the race. 3.2. experiment prague letna the monitoring of fans was realized at football stadium prague letna during the spring part of the czech football league (2014/2015). the stadium figure 3. construction of monitored grandstands: football stadium – prague letna(upper) [2], football stadium – prague eden (lower) [3]. is the home stadium of the team ac sparta praha. the team has one of the faithful and dangerous fans (figure 1) in the czech republic. ac sparta praha reached around second place of football table. the attendance of fans was falling down during the end of the football season. sparta finished at the second place of the football season. in the table 1, there are the described football matches and the occurrence of the synchronized jumping of the 2nd and the 3rd type (figuge 7). 3.3. experiment pilsen the experiment was realized at the ice-hockey stadium pilsen during all the czech ice-hockey league (2015/2016). long-term monitoring was cocerned to spectators´ behavivour during ice-hockey matches as for the football matches. specifically, the monitoring 153 martin verner acta polytechnica ctu proceedings figure 4. construction of monitored grandstands: ice-hockey stadium – pilsen (upper) [5], biathlon stadium – new town in moravia (lower) [4]. was focused on fans (figure 2) supporting team hc skoda plzen. hc skoda plzen reached upper part of the icehockey table. the team finished at the fourth place of the ice-hockey regular season. after regular season, pilsen team played ice-hockey play-off and it finished at the semi-final. the experiment recorded twenty games of the regular season of possible twenty-six and six games of play-off. the synchronized jumping of all rows did not occur when the camera was recording during the experiment. this behaviour is not a part of the ice-hockey fans style to support their team. this is a marginal issue of cheering. in 2016, the winter classic game (outdoor game) was played in brno. the match is a part of ice-hockey season between hc kometa brno and hc škoda pilsen. assembled grandstands (figure 5), similar to the one in new town in moraqvia, where build for this match. as the grandstands had similar construction, jumping was also prohibbited, but this ban was broken. the vibrations of grandstands could be felt, tribunes were very uncomfortable. figure 5. construction of grandstands in brno [6]. 3.4. experiment prague eden the monitoring of fans (figure 1) was realized at the football stadium prague eden again during the spring part of the czech football league (2015/2016). the stadium is the home stadium of the team sk slavie praha. the team is like ac sparta praha, it has one of the faithful and dangerous fans in the czech republic. sk slavie praha reached around the fifth place of the football table. attendance of fans was not falling down with the end football season. in the table 2, there are the described football matches and the occurrence of the synchronized jumping of the 2nd (figure 6) and the 3rd type. 4. conclusion the results of the experiments show the big differences between fans of behaviours at football stadium and at the other stadiums. behaviours of crowds were affected by the instructions from the organizers during the biathlon races. the synchronized jumping was not observed at any ice-hockey game during the experiment in pilsen. only the author of text was a direct witness of the synchronized jumping crowd during the winter classic game in brno. fans realized the synchronized jumping of all rows at least once at every football match during the experiment. difference of behaviour between stadium prague letna and stadium prague eden was the different used types of jumping. in the stadium prague eden, fans were preferring jumping all rows – horizontally, on the other hand the spectators were jumping all rows vertically in the stadium prague letna. the results of the experiment are the fundamentals for the next evolution. it describes the behaviour of the fans, the gained data will be determined further in the other research: how many people do it and how the attendance influences the level of the vibration of grandstands. acknowledgements this paper has been supported by the czech science foundation project no. ga15-15728s and faculty of civil engineering, czech technical university in prague (sgs project no. sgs16/201/ohk1/3t/11). 154 vol. 13/2017 the occurrence of synchronized jumping crowd figure 6. the sample of the grandstand vertical vibration (praha eden) – jumping rows – vertically [3]. figure 7. the sample of the grandstand horizontal vibration (praha letna) – jumping rows – horizontally [3]. date matchweek rival types of jumping jumping of the 2nd jumping of the 3rd march 8, 2015 19. fk dukla praha 2× 2× march 21, 2015 21. fk teplice 1× 1× april 26, 2015 25. fc zbrojovka brno 1× may 15, 2015 28. 1. fc slovacko 2× 1× may 30, 2015 30. fc bohemians praha 1× table 1. table of occurrence of synchronized jumping crowd during the match – ac sparta praha. date matchweek rival types of jumping jumping of the 2nd jumping of the 3rd february 28, 2016 19. fc banik ostrava 1× 1× march 11, 2016 21. fc fastav zlin 1× april 23, 2016 26. fk teplice 1× may 8, 2016 28. 1. fc slovacko 1× table 2. table of occurrence of synchronized jumping crowd during the match – sk slavie praha. 155 martin verner acta polytechnica ctu proceedings references [1] m. verner. real crowd behavior on the grandstands. 4th conference nano & macro mechanics 4(6):215–220, 2013. [2] m. verner, t. plachy, m. polák. an experiment focused on fans behaviour and induced grandstand vibrations during a football match. trans tech publications offices (4):75–78, 2016. [3] m. polák, t. plachy, m. verner. a comparison of spectators induced grandstand vibrations and fans behavior during two ice hockey games between the same teams in the regular season and in the playoffs. 3rd international conference on structural and physical aspects of construction engineering 3(4), 2016. [4] j. hubka. dynamic analysis of grandstand. master’s thesis. 2017. [5] m. verner. behaviour of fans during the football match. trans tech publications 102(4):77–80, 2016. [6] peri. winter classic lužánky, brno. https://www. peri.cz/projekty/kulturni-a-sportovni-akce/ winter-classic-luzanky.htmll. 156 https://www.peri.cz/projekty/kulturni-a-sportovni-akce/winter-classic-luzanky.htmll https://www.peri.cz/projekty/kulturni-a-sportovni-akce/winter-classic-luzanky.htmll https://www.peri.cz/projekty/kulturni-a-sportovni-akce/winter-classic-luzanky.htmll acta polytechnica ctu proceedings 13:153–157, 2017 1 introduction 2 types of fans behaviour 3 realized experiments 3.1 experiment new town in moravia 3.2 experiment prague letna 3.3 experiment pilsen 3.4 experiment prague eden 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.24.0031 acta polytechnica ctu proceedings 24:31–37, 2019 © czech technical university in prague, 2019 available online at https://ojs.cvut.cz/ojs/index.php/app effect of nacl on neutron flux density in graphite block tomáš slančík, kamil števanka∗ brno university of technology, faculty of electrical engineering and communication, technická 10, 616 00 brno, czech republic ∗ corresponding author: kamil.stevanka@vutbr.cz abstract. aim of this work is to study the effect of nacl on neutron flux through graphite block. graphite block had 39 channels drilled through it and filled with nacl. au and in foils and in wire was used for neutron activation measurement of the neutron flux. ambe neutron source with nominal activity of 92.5 gbq was used. the results were compared with mcnp simulations with channels filled with salt and air. twelve different locations for au and in foils were used and in wires were place in three locations to measure vertical profile of neutron flux. keywords: hpge, nacl, chloride salts, neutron flux. 1. introduction molten salts were first considered for application in nuclear reactors in late 1940s as part of the aircraft reactor experiment. molten salts were chosen because of the high power density requirements for the reactor and multiple other reasons [1, 2]. the research continued with molten salt reactor experiment (msre), which was considered a success [3, 4]. however the research of msrs was discontinued after msre due to other priorities, lack of funding and other political reasons [5]. since fluorides were chosen as primary coolant, most of the research to date focused on these mixtures [1, 3]. nevertheless chlorides were considered as well and therefore some cl based mixtures were studied too [2]. main advantages of msrs compared to pwrs are higher fuel burn-up, higher operating temperature and low operating pressure usually around 0.3 mpa. there is also theoretical possibility of breeding fuel from both uranium and thorium. among the challenges are higher neutron flux, corrosive behavior of salts, more difficult maintenance and inspection [6]. the interest in msrs was renewed in 21st century. even though most concepts are based on fluorides, some concepts, especially fast reactors, rely on chloride mixtures. among the concepts that plan to utilize chlorides are msfr by terrapower, llc and ssr by moltex energy ltd [7, 8]. apart from usage in msrs, chlorides can also be used in accelerator-driven systems [9]. since the data for neutronics of chlorides are not as well studied as fluorides, the goal of the work was to construct the experimental setup and determine how the nacl affects neutron flux density and to determine whether the experiment provides any useful data. this experiment was conceived to be the first in the series of experiments aimed at gathering more data about chloride neutronics. 2. experimental setup the equipment used for the experiment consisted of graphite block, ambe neutron source, sodium chloride and golden and indium threshold activation detectors (tad). dimensions of the graphite block were 506 × 300 × 300 mm (length x width x height). altogether 39 channels were drilled all the way through the block, these were filled with nacl. one hole was drilled at the center of the shorter side of the block to accommodate neutron source. this hole was drilled to the depth of 175 mm. the layout of the channels can be seen in figure 1. the diameter of the nacl channels was 29.3 mm and the diameter of neutron source hole was 22 mm. the block was placed on the 5 mm thick steel plate, which was attached to 600 mm tall wooden base. ambe neutron source with neutron emissivity of 5.6 · 106 s-1 was used. it is made of 833 mg of amo2 and 7374 mg of 9be encased in stainless steel cylinder which has diameter of 19.1 mm and height of 48.6 mm. 3. methodology preliminary calculations were made to determine suitable diameter. eventually it was decided to drill channels with diameter of 29 mm based on the practical feasibility. the 29 mm drill was the largest readily available and smaller diameters would mean more channels which would be time consuming. one channel took about 20 minutes to drill. due to the concerns about graphite fragility it was decided that distance between edge of the holes and the block should be at least 15 mm and keeping sufficient distance between edges of holes meant that there would be actually less volume of salt if bigger diameter was used. the channels are 0.3 mm wider due to drill vibrations. the hole for neutron source was drilled to the 31 https://doi.org/10.14311/app.2019.24.0031 https://ojs.cvut.cz/ojs/index.php/app tomáš slančík, kamil števanka acta polytechnica ctu proceedings figure 1. layout of drilled channels. depth of 175 mm in order to vertically center the source. solidworks 2014 software was used to create model of the block and the base. each of the 29 mm channels was filled with 99.5 % purity nacl salt, the salt was put into polyethylene bag for easier manipulation. neutron activation measurement was used to measure the neutron flux density. 197au (figure 4) and 115in (figure 3) foils were used as activation detectors, which have high cross section for thermal neutron capture. foils were placed in paper wrap in the selected channels as seen in figure 2. foils were placed at the 3 o’clock side of the holes, approximate placement of the foils is designated by red dots in figure 2. all of the in foils had same diameter of 16 mm, thickness of 0.5 mm and 99.99 % purity. au foils had square shapes but their dimensions differed as can be seen in table 1. apart from the foils, in wires were used in channels iiα, iiγ and iiη, right in front of the neutron source. this was used to determine neutron flux over the entire height of the block. the diameter of in wire was 1 mm and length was 300 mm. the au detectors were placed lower than 150 mm from the base and in detectors were placed above 150 mm because the foils could not be placed side by side in the vertical center of the channels at height of 150 mm. in addition to the detectors in the channels, other detectors were placed on the external sides of the block. both au and in detectors were placed in the center of the respective sides of the block. detectors with cadmium shielding were added to the left and rear sides (from the source point of view) of the block to detect ratio between thermal and fast neutrons. in hole purity(%) thickness (mm) dimensions (mm) iiα 99.95 0.25 10x10 iiγ 99.95 0.25 10x10 iiη 99.90 0.02 13x13 3a 99.95 0.25 10x10 3c 99.90 0.02 13x13 3e 99.90 0.02 12x12 iiiα 99.95 0.25 10x10 iiiγ 99.90 0.02 12x12 iiiη 99.90 0.02 13x13 4a 99.90 0.02 13x13 4c 99.90 0.02 12x12 4e 99.90 0.02 12x12 table 1. au detectors characteristics. foils had same dimensions as the ones in the channels and au foils are described in table 2. hpge detector was used to measure activation of the foils. gamwin software was used to evaluate measurements. the reaction rate rrc (as seen in equation 1) was calculated after the area of measured peaks was manually determined using results from gamwin software. intensities and detector efficiency for designated gamma energies can be seen in table 3. 32 vol. 24/2019 effect of nacl on neutron flux density in graphite block figure 2. layout of detectors. figure 3. indium-115 radiation capture cross-section [10]. figure 4. au-197 radiation capture cross-section [10]. 33 tomáš slančík, kamil števanka acta polytechnica ctu proceedings side purity(%) thickness (mm) dimensions (mm) source 99.95 0.250 10x10 left cd 99.95 0.250 25x25 left 99.95 0.125 25x25 right 99.95 0.125 25x25 rear cd 99.95 0.500 25x25 rear 99.95 0.500 25x25 table 2. outside au detectors characteristics. material eγ (kev) iγ (%) �γ (-) in 416.860 27.7 0.03484 1097.326 56.2 0.00833 1293.558 84.4 0.00559 au 411.802 96 0.03525 table 3. measured energies and characteristics. rrc = s(eγ) · λ · trealtlive �γ(eγ) · iγ(eγ) · (1 − e−λtirr )· · (1 − e−λtreal ) · e−λtdelay (s−1) (1) where s(eγ) is area of measured peak of selected energy level, λ is decay constant of particular isotope, treal is measuring time, tlive live time of measurement, �γ(eγ) is detector efficiency for given energy level, iγ(eγ) is intensity of gamma line, tirr is duration of irradiation and tdelay is time between the end of irradiation and beginning of measurement. afterward the weighted average of reaction rates for different energies for in was calculated. then reaction rates per gram of material was calculated using dimensions of detectors and densities of materials [11],[12],[13]. the irradiation lasted for 7 days and measuring intervals were 15 mins for in wire and 20 mins for in and au foils. the irradiation lasted for 7 days to ensure sufficient degree of activation because half-life of 198au is 2.694 days [14]. because samples could not be reinserted into the setup without disrupting the experiment, it was not possible to conduct multiple measurements of the samples. reaction rates measured and calculated on outside detectors with and without cadmium shielding were used to calculate cadmium ratio 2 rcd = rr rcdr (2) where rr is reaction rate of all of the neutrons, rcdr is reaction rate of fast neutrons figure 5. vised visualisation. 3.1. mcnp simulation mcnp 6 was used to simulate the experiment [15]. polyethylene was neglected in the model given that combined weight of the bags was approximately 100 g compared to the weight of the entire setup which was approximately 90 kg. the model contained additional 300 mm thick wall on the side of neutron source, this could have some impact on the results. activation detectors were not placed within the channels but simplified approach was taken which evaluated neutron flux density through designated sections of channels. channels were separated into four sections, the two evaluated sections were at heights between 140 to 150 mm and 150 160 mm. then the reaction rate was calculated using calculated neutron fluxes, atomic density and cross sections of au and in. simulations were done for holes filled with air and nacl. 4. results the results of experiment and simulations can be seen in graphs figure 6 to 13. some results have substantial error margins, this is due to low intensity of neutron flux of neutron source and relatively short time of measurements. the short measuring time was caused by high number of samples because there was only one hpge available. the measuring time could not be longer, because the activity of samples would decrease too much and the setup did not allow to reinsert foils. cadmium ratio in table 4 shows that neutrons were not very well moderated, for future research would be useful to have higher volume of graphite. as can be seen in the graphs figure: 6 to 9, the trend of in foils is similar to mcnp simulation results, however most of the reaction rates are higher in the experimental samples than in simulations. this could be due to aforementioned simplifications and also due to high degree of uncertainties. the result for in foil in position αiii (figure 8) is different, but the error margin for that sample is also substantial, therefore it cannot be concluded that the model is inconsistent with the experiment. the results for golden foils of 0.02 mm thickness differ substantially from simulations because reference reaction rate was calculated for thickness of 0.25 mm in mcnp simulations. the experimental results for 34 vol. 24/2019 effect of nacl on neutron flux density in graphite block α γ ε location 0 20 40 60 80 100 r r c ( − 1 g − 1 ) reaction rate of indium in channel in po ition ii mcnp-salt mcnp-air experiment figure 6. results for in foils in positions ii. 0.25 mm thickness is much more similar to the simulation, whereas reaction rate for 0.02 mm is much higher, this is caused by self-shielding of the thicker foil material. the difference between simulations with salt and air is not very pronounced. this is due to relatively small amount of salt in the system, however all simulations show slightly higher reaction rates for air in the farthest positions, indicating that nacl captures small amount of the neutrons. wire measurements figure 14, figure 15, figure 16 showed high degree of uncertainty because of short half-life of in wire and short measuring time. each wire was divided into five segments, each 6 cm long. the wires could not be reinserted, therefore they were cut into 5 pieces due to short half-life [14] and these segments were measured. the first wire showed expected results with higher rate in the middle, other two wires showed different results with no clear trend. this was most likely caused by short measuring time, low activity of the samples farther from the neutron source and short half-life of indium. the results for in wire are not reliable and require further measurements. position indium goldrcd (-) rcd (-) left 1.4090 1.5579 rear 1.3243 1.6416 table 4. experimental results. 5. conclusions the paper studies neutron flux density within graphite block with drilled channels that were filled with nacl. the aim of the work was to design and construct experimental setup and get some experience with chloride salts. another goal was to see whether the mcnp α γ ε location 20 40 60 80 100 r r c ( s( 1 g − 1 ) reaction ates of indium in channels in position 3 mcnp-salt mcnp-air experiment figure 7. results for in foils in positions 3. α γ ε location 10 20 30 40 50 60 r r c ( s( 1 g − 1 ) reaction ates of indium in channels in position iii mcnp-salt mcnp-air experiment figure 8. results for in foils in positions iii. model is reasonably accurate compared to the experiment and to find out ways in which we could improve the experimental setup. the main purpose of the setup is to provide students with opportunity to measure and evaluate real life data and get hands on experience with scientific work. neutron activation measurement was used to determine neutron flux. twelve au and in foils were used and three 30 cm long in wire were used as well. the wire measurement had high error margin due to short measuring time and only the wire at the closest position to the neutron source showed expected trend with regard to vertical distribution of neutron flux. both in and au foils showed similar trends to the simulation, reaction rates were higher than those of the simulations in all cases except one and this could be attributed to measuring error. the self-shielding effect of gold affected the results significantly and measured values for thinner, 0.02 mm thick foils diverged from simulation considerably. this was caused by simulation input which used thicker 0.25 mm thick foils for normalisation. despite the fact that 0.02 mm foils showed significant divergence in values compared to model, it nevertheless showed similar trends. given the error 35 tomáš slančík, kamil števanka acta polytechnica ctu proceedings a c e location 10 20 30 40 50 r r c ( s− 1 g − 1 ) reaction ates of indium in channels in position 4 mcnp-salt mcnp-air experiment figure 9. results for in foils in positions 4. α γ ε location 10 20 30 40 r r c ( − 1 g − 1 ) reaction rate of gold in channel in po ition ii mcnp-salt mcnp-air experiment figure 10. results for au foils in positions ii. margins it is sufficient indication that the model is in good agreement with the experiment. the measuring errors were quite large, this was due to limitations of experimental setup which did not allow for reinsertion of the foils. this was particularly visible on in wires which were separated to 5 segments, but because of short half-life of indium could not be measured for a long time. in the future it is planned to upgrade the setup to allow for better manipulation of salt and also reinsertion of detectors. the samples could not be reinserted because they were inserted simultaneously with the polyethylene bags. once they were removed the polyethylene bag would not allow for the sample to be put back in. in the future it is planned to use vessels that could keep their shape, probably aluminium tubes. thus allowing for continuous measurement. after the upgrade of the experimental setup it is planned to repeat the experiment, perform another experiment with empty channels to compare the results and adjust the mcnp model. it is also planned to measure the effects of other chloride salts, such as licl, kcl, cacl2, mgcl2. it is also possible to add another layer, approximately 200 mm of graphite between neutron source and the block to get higher amount of thermal neutrons. a c e location 10 20 30 40 50 60 70 80 r r c ( s− 1 g − 1 ) reactio rates of gold i cha els i positio 3 mcnp-salt mcnp-air experiment figure 11. results for au foils in positions 3. α γ ε location 10 20 30 40 50 60 r r c ( s( 1 g − 1 ) reaction rates of gold in channels in osition iii mcnp-salt mcnp-air experiment figure 12. results for au foils in positions iii. acknowledgements the research was supported by project ltausa18198 and is part of cooperation between but and texas a&m university. authors are thankful to their supervisor and pi of the project, karel katovsky, ph.d., for his support and advice during the project. references [1] w. cottrell, h. hungerford, j. leslie, j. meem. operation of the aircraft reactor experiment. ornl -1845, oak ridge national laboratory, 1955. http://large.stanford.edu/courses/2016/ph241/ dodaro1/docs/ornl-1845.pdf. [2] d. f. williams, k. t. clarno. evaluation of salt coolants for reactor applications. nuclear technology 163(3):330–343, 2008. doi:10.13182/nt08-a3992. [3] r. robertson. msre design and operations report description of reactor design. ornl-tm-728, oak ridge national laboratory, 1965. http://moltensalt.org/references/static/ downloads/pdf/ornl-tm-0728.pdf. [4] m. rosenthal, p. kasten, r. briggs. molten-salt reactors history, status, and potential. nuclear appliactions & technology 1969. http://moltensalt.org/references/static/ downloads/pdf/nat_msrintro.pdf. 36 http://large.stanford.edu/courses/2016/ph241/dodaro1/docs/ornl-1845.pdf http://large.stanford.edu/courses/2016/ph241/dodaro1/docs/ornl-1845.pdf https://doi.org/10.13182/nt08-a3992 http://moltensalt.org/references/static/downloads/pdf/ornl-tm-0728.pdf http://moltensalt.org/references/static/downloads/pdf/ornl-tm-0728.pdf http://moltensalt.org/references/static/downloads/pdf/nat_msrintro.pdf http://moltensalt.org/references/static/downloads/pdf/nat_msrintro.pdf vol. 24/2019 effect of nacl on neutron flux density in graphite block a c e location 10 20 30 40 50 60 r r c ( s− 1 g − 1 ) reaction rates of gold in channels in osition 4 ex eriment mcnp-salt mcnp-air figure 13. results for au foils in positions 4. 0-6 6-12 12-18 18-24 24-30 part of the wire (cm) 40 60 80 100 120 r r c ( s− 1 g − 1 ) reaction rates of indium wires in position iiα figure 14. results for in wire in position α. [5] b. hoglund. why the molten salt reactor (msr) was not developed by the usa. blogpost, 2010. http: //moltensalt.org/references/static/downloads/ pdf/whymsrsabandonedornlweinbergsfiringv3.pdf. [6] d. e. holcomb. molten salt reactors today status & challenges. in workshop on msr technologies commemorating the 50th anniversary of the startup of the msre. ornl, 2015. https://public.ornl.gov/conferences/msr2015/ pdf/05-current%20msr%20status%20and% 20challenges_david%20holcomb.pdf. [7] world nuclear association molten salt reactors, 2018. http: //www.world-nuclear.org/information-library/ current-and-future-generation/ molten-salt-reactors.aspx. [8] m. elsa, h. daniel, a. m., et al. preliminary design assessment of the molten salt fast reactor. in european nuclear conference enc2012. 2012. [9] p. mcintyre, s. assadi, k. badgley, et al. acceleratordriven subcritical fission in molten salt core: closing the nuclear fuel cycle for green nuclear energy. in 22. international conference on application of accelerators in research and industry. aip, 2013. doi:10.1063/1.4802405. 0-6 6-12 12-18 18-24 24-30 part of the wire (cm) 20 30 40 50 60 70 r r c ( s− 1 g − 1 ) reaction rates of indium ires in position iiγ figure 15. results for in wire in position γ. 0-6 6-12 12-18 18-24 24-30 part of the wire (cm) 10 20 30 40 50 60 70 r r c ( s− 1 g − 1 ) reaction rates of indium ires in position iiε figure 16. results for in wire in position η. [10] nuclear energy agency java-based nuclear data information system, 2013. https://www.oecd-nea.org/janis/. [11] the engineering toolbox densities of common solids, 2019. https://www.engineeringtoolbox.com/ density-solids-d_1265.html. [12] the engineering toolbox metals and alloys densities, 2019. https://www.engineeringtoolbox. com/metal-alloys-densities-d_50.html. [13] k. katovsky. studium sekundarnich neutronu a jader vznikajicich pri reakcich protonu a neutronu v tercich z uranu a plutonia. ph.d. thesis, [czech], cvut, 2008. [14] nds iaea live chart of nuclides, 2019. https://www-nds.iaea.org/relnsd/vcharthtml/ vcharthtml.html. [15] c. werner. mcnp users manual code version 6.2, 2017. https://mcnp.lanl.gov/mcnp_manual.shtml. 37 http://moltensalt.org/references/static/downloads/pdf/whymsrsabandonedornlweinbergsfiringv3.pdf http://moltensalt.org/references/static/downloads/pdf/whymsrsabandonedornlweinbergsfiringv3.pdf http://moltensalt.org/references/static/downloads/pdf/whymsrsabandonedornlweinbergsfiringv3.pdf https://public.ornl.gov/conferences/msr2015/pdf/05-current%20msr%20status%20and%20challenges_david%20holcomb.pdf https://public.ornl.gov/conferences/msr2015/pdf/05-current%20msr%20status%20and%20challenges_david%20holcomb.pdf https://public.ornl.gov/conferences/msr2015/pdf/05-current%20msr%20status%20and%20challenges_david%20holcomb.pdf http://www.world-nuclear.org/information-library/current-and-future-generation/molten-salt-reactors.aspx http://www.world-nuclear.org/information-library/current-and-future-generation/molten-salt-reactors.aspx http://www.world-nuclear.org/information-library/current-and-future-generation/molten-salt-reactors.aspx http://www.world-nuclear.org/information-library/current-and-future-generation/molten-salt-reactors.aspx https://doi.org/10.1063/1.4802405 https://www.oecd-nea.org/janis/ https://www.engineeringtoolbox.com/density-solids-d_1265.html https://www.engineeringtoolbox.com/density-solids-d_1265.html https://www.engineeringtoolbox.com/metal-alloys-densities-d_50.html https://www.engineeringtoolbox.com/metal-alloys-densities-d_50.html https://www-nds.iaea.org/relnsd/vcharthtml/vcharthtml.html https://www-nds.iaea.org/relnsd/vcharthtml/vcharthtml.html https://mcnp.lanl.gov/mcnp_manual.shtml acta polytechnica ctu proceedings 24:1–7, 2019 1 introduction 2 experimental setup 3 methodology 3.1 mcnp simulation 4 results 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.23.0049 acta polytechnica ctu proceedings 23:49–53, 2019 © czech technical university in prague, 2019 available online at https://ojs.cvut.cz/ojs/index.php/app a retaining wall reconstruction in kutná hora jiří svoboda∗, dana hadačová pragoprojekt, a. s. (plc.), k ryšánce 1668/16, 147 54 prague, czech republic ∗ corresponding author: jiri.svoboda@pragoprojekt.cz abstract. this paper addresses the reconstruction of a historical stone retaining wall at the busy kremnická – táborská road in the town of kutná hora. the structure was in very poor condition. there was the danger of an accident involving a wall slide, or even a total collapse of the road structure. the construction site in question is located near the st. barbara’s church, in the area of the formerly important but now historical mine osel. with respect to the possibility of undermining, various geodetic surveys and observations of the surrounding area and rock mass were performed. keywords: retaining wall, undermined area, st. barbara’s church, kutná hora. 1. introduction on the basis of expert static assessment and juridical assessment of the retaining wall in kremnická and táborská streets a preliminary permit was issued in 2015 by the special building authority for the reconstruction works on the road no. iii/3377 pertaining to the condition of the current retaining wall and the adjacent part of the road. the wall was in serious disrepair and it was necessary for it to be fixed immediately, possibly because of the hazard of accident such as a wall slide, or even a total collapse of the road embankment. in the first stage of construction the necessary securing works were carried out in order to avert an emergency or to eliminate its possible effects and thus preventing any danger to health or life regarding the public. the aim of the following second stage of the rehabilitation works was to finish the whole wall and road structure and ensure long-term durability, safety and functionality. the construction site is near st. barbara’s church in the location of the important historical mine osel. with regard to the possibility of undermining, various geodetic surveys and observations of the surrounding area and rock mass were performed. a detailed research of historical mining materials was also done with the cooperation of local citizens. on the basis of output evaluations a definitive technical solution was adopted, mainly in areas of cave-ins, large underground space rehabilitation and roadbed remediation. 2. sources and geology the sources and inputs for design work were: • design documentation for building permit by top con servis, s. r. o. (ltd.), 03/2015, • engineering-geological survey by forvia cz s. r. o. (ltd.), 03/2015, • static assessment of the retaining wall condition by libor pokorný (authorized engineer for bridges and engineering structures), 05/2015, • construction site inspection done by the designer on 01/07/2015, which declared that the condition of the wall endangered traffic and therefore safety measures had to be taken immediately – and the very same day they actually were. on the basis of the above mentioned sources it was concluded that the major part of the retaining wall was in serious disrepair and the work for its securing and subsequent reconstruction were to start immediately. from the geomorphological point of view the location is part of the kutná hora plateau, which is formed by crystalline rock. the plateau surface is subject to intrusion by the steeply carved valleys of local watercourses. watercourse erosion has had a predominant impact on the formation of local steep rocky banks with bare sandstone strata benches. because of gravity impact and erosion processes, horizontally deposited sandstones split off along the predisposed discontinuity planes and slid into the valley in the form of blocks and benches of various sizes. the phenomenon of open crack occurrence and the detached blocks slight inclination to the valley and the watercourse is important for stability assessment on the edge of the valley slope. this gives a very uneven shape to the rock base in the wall foundation joint. the formerly important silver mine osel is located in the area concerned as is the osel vein range. in the middle ages the mining works were done under the sandstone layers in a zone of clay rocks or geological faults. the foundations of these workings were inadequately carried out, and that is why after the drifts were exhausted and dismantled, the clay rock gradually reduced to mush causing the present cave-ins. we faced this reality several times during the execution of drilling work when the drilling rig slumped. 49 https://doi.org/10.14311/app.2019.23.0049 https://ojs.cvut.cz/ojs/index.php/app jiří svoboda, dana hadačová acta polytechnica ctu proceedings anchors bond lengths were also located in the mushy surrounding because of the ancient mining. that led to a considerable overconsumption of the injecting material during the anchor bond lengths execution. for illustration, the underground scheme is shown in figure 1. the view of the original and new wall near st. barbara’s dome is captured in figure 2. the detail of the wall is depicted in figure 3. 3. essential requirements for the new design in designing the retaining wall reconstruction, the following requirements had to be taken into account: • two-way traffic in at least one lane had to be maintained in kremnická street during the excavations. • proper drainage of the retaining wall back. underground water was not expected, but there was a risk of rain water accumulation (including water from the road surface), because of leaking at the wall back, which might cause degradation of the bedrock. • the underground services and the surrounding buildings including the retaining walls on the opposite side of the street should not be endangered by the excavations. • the aesthetic design had to be made according to the requirements of the national heritage institute. 4. conception of the new design after meetings on conceptualisation with the ordering party (central bohemia road agency – ksus and kutná hora municipality) a basic variant of the retaining wall design was set up. a monolithic reinforced concrete cantilever wall with stone lining was designed to be built in a temporarily supported excavation [1]. the temporary excavation was supported by anchored piling in such a way that any interference with the existing underground services was excluded. the continuous raised guiding wall for embedding micropiles prevented car slide to the excavation, and therefore ensured some traffic safety in kremnická street and also kept rainwater out of the excavation. the reinforced concrete cantilever wall itself has a variable height circa 1.5–5.0 m. its total length is circa 315.0 m. it is founded on an underlay concrete slab 0.15 m thick. the breadth of the footing is within 1.9– 2.5 m and the height 0.45–0.65 m. the wall was mainly built in 5–6 m long sections. in some sections, where the geology was unfavourable or where there were additional static reasons, the footing was founded on steel micropiles. the wall back drainage is provided by passages embedded to the stem and the longitudinal drainage system behind the wall. a small stone wall 0.4 m thick and 0.9 m high was built over the upper level of the retaining wall. with respect to the location (barrier curb, design velocity) there was no need to give this stone wall some dimensions as a safety barrier for traffic. the face of the new retaining wall is also walled in with stone. this lining is in the upper part gradually merged with the stone wall structure and capped with a flat stone roof 0.55 m wide. the aboveground stone part and the stone lining are integral parts of the retaining wall static system. local and overall stability of the retaining wall structure was assessed in its particular characteristic sections. 5. geotechnical monitoring and monitoring board geotechnical monitoring (gtm) was carried out for the contract owner directly. gtm is mandatory for structures where serious negative impacts are possible for the public or either built or neighbouring objects. the placing of some measuring points including the survey of them was undertaken in advance of the actual construction work. the rules for implementation of gtm and monitoring board leadership were given in the tender bid documentation. detailed passport of all affected buildings and underground services was made. for the second stage a new contractor was chosen. that meant, a re-passport of all objects had to be made in the zone of influence and at the same time all the observations and monitoring within the gtm had to be evaluated. the second re-passport was made at the end of construction, resulting in only two objects being damaged in the context of the construction. a geodetic survey on objects and walls, inclinometrical observation of the rock mass, crack monitoring on objects and also regular visual control of influenced objects and underground services were all done in the field. the outputs were regularly evaluated and construction procedure was continuously modified according to the monitoring board and its expert conclusions. simultaneously, the geological condition found on site was monitored with expert specialist geotechnical supervision. on the basis of output evaluations the definitive technical design was modified, mainly in the area of cave-ins and large underground space rehabilitation. subsequently static calculations had to be resolved with impacts on design detail. two major cave-ins happened on the site during the construction. right under the carriageway a free underground space of circa 65 m3 was rehabilitated. some other surveying drills were made to find other free spaces especially under the carriageway. all these caves were filled with cement-ash mixture followed by a low-pressure cement injection. completion of the rehabilitation of surrounding buildings, started in advance, was also finished within the monitoring. 50 vol. 23/2019 a retaining wall reconstruction in kutná hora figure 1. underground space scheme. 6. on the history of mining in kutná hora on the basis of knowledge gathered from local citizens during re-passport and research of archival materials, especially the monograph "kutnohorské dolování" by j. bílek [2], it could be ascertained that the site is in the area of the important ancient mine osel (named by a burgomaster called asinus) in the osel vein range. at least 50 bigger mines are estimated to have been operated in the town area from the 13th to the 16th century. the majority faded out during the hussite wars. the osel mine manifested the wealth and heroic status of the miners in the medieval town of kutná hora. this mine became a status symbol of the miner and therefore a focus for mining authority. not just the beautiful and impressive churches of st. barbara and st. jacob (in kutná hora) were built from the mining profits, but also contributions had to be made to st. vitus cathedral in prague. mining in the osel range expanded in the 13th to 15th centuries and it penetrated to the depth of almost 500 m. in the year 1541 an underground water outflow to osel took the life of eight miners and signalled the end of mining in 1545. the main shaft location has not been found until now. it is estimated to be located near the jesuit college. the other shafts were located close to each other nearby the limestone quarry behind st. barbara’s (mladé and staré rousy). the waste bank is documented in the area of žižkov and near všech svatých. the deepest point of this complex was located under the zelená hora street loop above the bylanka valley. the oldest shafts in the area concerned were opened probably before the end of the 13th century, when the upper parts of the ore veins were mined. these were settled within small horizontal distances of 20 to 50 m with cross-sectional dimensions of 1.2×1.5 m, or 1.5×2.0 m. in loess and marl the shafts were timbered. the mines were drained by drainage galleries driven from the vrchlice valley. these galleries hooked up to the shafts at a level of 30 to 50 m under the surface. shaft backfills were built almost entirely from the slag heaps. over centuries they were affected by underground water circulation. naturally, that would be the reason for the gradual relaxation and also often for the cave-ins. additionally, in the shafts the clay rock gradually reduced to mush, again causing cave-ins. the risk of cave-ins in these ancient shafts and galleries is nowadays small, although it cannot be completely excluded. therefore, any possibility must be taken into account for obvious safety reasons. the whole osel range consists of five rich ore veins stretching along the left side of the vrchlice river from holy trinity church to the western part of the town. the very old mining workings stretch from bylany over so-called cech and vyšata quarry to the west of st. barbara’s dome. indeed that is why the relics of mining were expected to be found on the construction site and that was proved by object movements in both táborská and kremnická streets. these movements were not caused by the security actions for construction but by the presence of these very old and defunct mine workings. 51 jiří svoboda, dana hadačová acta polytechnica ctu proceedings figure 2. view of the original (top) and new wall (bottom) near st. barbara’s dome. 7. georadar – a geophysical survey method with respect to the above mentioned facts the question of other disrupted places was raised during the construction process. these faults could in the future endanger not only the new retaining wall but also the adjacent carriageway that has itself been reconstructed during the construction including the new structural pavement layers. gravimetric data were processed and evaluated. the principal result of the survey led to the finding that there was several other places under the road where there existed relatively continuous ranges with a lack of material. from the geophysical point of view it was recommended to rehabilitate these places preferably with injections to the depth of 10 m. the use of georadar in the final stage of the construction process in order to check the roadbed proved to be positive and valid [3]. it was decided to replace the original pavement design (reduced pavement structure in one half of the carriageway and only reconstruction of asphalt layers in the other half) for a complete reconstruction of the road structure in the whole rebuilt section. 8. conclusions design work started with the contract owner’s request to appraise the static assessment for the original retaining wall by libor pokorný. just after the first inspection of the wall and road in question the author of this appraisal pronounced the static situation to be very serious, and endangering not only traffic. the issue gradually developed from a small problem of carriageway slump and collapse of a rather small part of the wall to the bigger problem of global wall instability including the presence of undermining under the road. the amount of work needed gradually increased. the situation had to be solved quickly and as a complex whole. the designer’s thoroughness and often indeed intransigence paid off because a complex interconnected solution was carried right through to finality. the active cooperation of all affected authorities and organizations, contract owners – ksus and the municipality of kutná hora, with both contractors and the designer was a necessary condition for managing the situation. property relations responsibilities complicated the problem solving, because the road owner is the central bohemian region, while sidewalks and some of the walls belong to the town of kutná hora, and then leaving a major part of the walls belonging to private owners. for this reason a memorandum for a common course of action was initiated. besides the property municipality department and traffic municipality department, the officers from municipality department for national heritage, education and culture of kutná hora took part in the monitoring board. they consulted each other actively regarding the shape and appearance of the new wall and participated on the selection of proper stone, or possibly and more precisely the source quarry for the new wall lining. respected experts from the department of geotechnics of the fce of ctu prague (prof. jiří barták and assoc. prof. jan masopust) also made a considerable contribution. taking account of the danger of delay a very unusual instance of cooperation with the private owners of the surrounding estates was established, and this led to a prompt start to the process of implementing the geotechnical monitoring and then on to the construction work itself. the structure was deemed in the 3rd geotechnical category because of its complexity, and this made gtm necessary. local citizens themselves understood the importance of the situation and enabled their objects for passport and surveying point installation and facilitated incident-free surveys. the initial meeting was held in the municipal hall with the attendance of the town officials, representatives of the offering party, contractor and designer. the designer gave a detailed presentation on the design construction activities. an individual approach to the private owners in the affected zone proved to be crucial. personal exchange and discussion on any particular problems 52 vol. 23/2019 a retaining wall reconstruction in kutná hora proved much more advantageous than relying on written receipted and recorded correspondence. it brought benefits during the construction process, when hazardous works were performed near the objects and there was a need for safety monitoring and controlling the condition of objects. mutual reliance among all partners in the building process, involving the fulfilment of assurances and keeping to deadlines, proved to be an essential condition for successful construction management of such a complex structure. references [1] p. turček. foundation engineering (zakládání staveb). jaga, bratislava, 2005. [2] j. bílek. kutnohorské dolování. kuttna, 2000. [3] j.-l. briaud. geotechnical engineering: unsaturated and saturated soils. wiley, new york, 2013. figure 3. structural condition comparison: the old (top) and the new wall (bottom). 53 acta polytechnica ctu proceedings 23:44–48, 2019 1 introduction 2 sources and geology 3 essential requirements for the new design 4 conception of the new design 5 geotechnical monitoring and monitoring board 6 on the history of mining in kutná hora 7 georadar – a geophysical survey method 8 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2019.25.0032 acta polytechnica ctu proceedings 25:32–35, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app direct measurement of reaction forces during fast dynamic loading jan falta∗, marcel adorna, tomáš fíla, petr zlámal czech technical university in prague, faculty of transportation sciences, department of mechanics and materials, konviktská 20, 120 00 prague 1, czech republic ∗ corresponding author: falta@fd.cvut.cz abstract. the presented paper is focused on embedding of the serially manufactured piezo-electric impact loadcell into open hopkinson pressure bar (ohpb) and it’s modifications for a direct force measurement during dynamic loading. conventionally, during the shpb test dynamic force equilibrium is investigated by a comparison of the transmitted signal wave and the difference between the incident, reflected and transmitted signals measured by strain gauges [1]. however, in the experiments with specimens with low mechanical impedance, a major portion of the incident wave is reflected back on the interface between the bar and the specimen. comparison between two-large amplitude incident and reflected pulse and a small-amplitude transmitted pulse can be influenced by large error and resulting force equilibrium can be inaccurate. therefore, a piezo-electric quartz impact force transducer was used to directly measure the axial forces in the vicinity of the specimen end surfaces. keywords: hopkinson bar, instrumentation, strain-gauge measurement, piezo-electric loadcell. 1. introduction over the past decades, dynamic mechanical testing become widely used method for determination of properties of modern materials (e. g. metal foams, laser sintered structures). reliable recording of the dynamic response of the material during impact testing is therefore essential for obtaining required material properties at high strain rates. the strain gauges are one of the most commonly used sensors in dynamic testing devices (e. g. split hopkinson pressure bar, charpy impact test, drop tower). another way to measure reaction forces can be use of piezo-electric loadcell. in order to obtain reliable data, these sensors must be properly calibrated in a expected range of measured forces starting at 2 kn up to 20 kn. according to fujii, et. al., 2003 [2], there are three methods for calibrating dynamic force transducers. calibrating against oscillation forces, impact force or against a step force. however, all of these methods require expensive equipment and a calibrated reference sensor. this paper presents embedding of the serial manufactured piezoelectric impact load cell into ohpb (open hopkinson pressure bar) and its calibration using two pairs of strain gauges as reference sensors. calibration of the strain gauges can be done using static loading procedure without much difficulty was demonstrated in the previous research [3]. 2. testing setup ohpb, modification of standard shpb apparatus was used for calibration of the force transducer. in this case incident bar was guided by a linear guidance system and was instrumented with strain gauges (see figure 1). the incident bar with length of 1600 mm and diameter of 20 mm made from high strength aluminum alloy (en-aw-7075) was used. the impact bar was fitted with foil strain gauges 3/120 ly61 (hbm, germany) with active length 3 mm wired in the wheatstone half-bridge arrangement at a distance of 200 mm from the face of the bar. the nominal resistance deviation of the strain gauge according to manufacturer is 1 %. the transmission part consists of two bars of the same material with lengths of 230 mm (first bar) and 1470 mm (second bar) between which a load cell is mounted. these rods are supported by a low friction polymer housings (drylin series, igus, germany) and instrumented using same strain gauges as on the incident bar. calibration of the strain gauges was carried out by gradual increasing of static load by 100 n step. both strain gauges are located at a distance of 60 mm from the load cell. since the load cell is not designed to measure tensile forces, and in the standard configuration the tensile wave is reflected at the free end of the transmission bar, a third rod was added to avoid the tensile wave from returning to the load cell and preventing cause possible damage. instrumentation of the load cell and strain gauges can be seen in figure 2. 3. piezoelectric force transducer most of the dynamic force transducers are based on piezoelectric crystals. principle of piezoelectric crystals stands on the generation of electric charge due to compression by the external force [4]. thus, the generated electric charge is converted as a small change in voltage at the output and readout by a corresponding measuring device. the main reason for using of piezo32 http://dx.doi.org/10.14311/app.2019.25.0032 http://ojs.cvut.cz/ojs/index.php/app vol. 25/2019 direct measurement of reaction forces figure 1. ohpb setup scheme figure 2. experimental setup overview electric crystals for dynamic loading measurements is their high natural frequency limit (commonly in order of 100 khz) and quasi-linear behavior over a wide amplitude range (pcb piezotronics, 2012) [5]. however, these sensors are unable to measure static and quasistatic loads or tensile load without prestressing. in this paper, a quartz crystal force sensor type 200c20 (pcb piezotronics, inc.usa) was used in calibration experiments. this sensor is designed to measure compression forces in a range starting at units of n up to 88.96 kn with sensitivity of 56.2 mv/kn and upper frequency limit of 40 khz [6]. this type of sensors in general have quasi-linear output above approx. 20 % of full scale output (fso) (see figure 3). figure 3. output applied force [5] if the measured forces are lower than this threshold, it is necessary to preload the sensor (quartz) to shift into the linear region. in our case the dimension of thread in the load cell and both bars is 1/4"-28 with tensile strength approx. 10 kn for steel thread and bolt. the 20 % load threshold for instrumented load cell is 17.8 kn. preload above this threshold would require the use of a strength screw of class 10.9. or 12.9 which significantly exceeds the strength capacity of the alluminium rod. therefore, it was not possible to applied enough preload force to reach linear output dependency. 3.1. load cell calibration force sensor was mounted between two bars with steel bolt using a threaded steel bolt (see figure 4). two strain gauges were instrumented on adjacent bars in the distance of 60 mm (see figure 1). for the load cell figure 4. quartz force sensor arrangement [6] 33 j. falta, m. adorna, t. fíla, p. zlámal acta polytechnica ctu proceedings calibration, 38 void tests (direct impact of the incident bar without sample) were performed in load range of 2 kn 28 kn. the signal measured by the load cell was converted to the unit of force using sensitivity constant (17.241 n/mv) given by the manufacturer. the force signal measured by the load cell was compared with signals from the strain gauges and a growing deviation was observed with decreasing applied force. selected waveforms of the measured signals for three different loads of 2 kn, 8 kn and 15 kn are shown in figures 5 7. the correction coefficient was determined as the ratio between the mean value of forces measured by strain gauges and average force measured by the load cell from the plateau area of the deformation pulse. it can be seen that the shape of the load cell signal corresponds with the strain gauges signal and the difference between the load cell and strain-gauge is decreasing with the applied force. the overall signal-tonoise ratio with used instrumentation is significantly lower for load cell signal. figure 5. calibration experiment at low impact force (approx. 2 kn) with deviation corresponding to force ratio coefficient of 1.315 figure 6. calibration experiment at low impact force (approx. 8 kn) with deviation corresponding to force ratio coefficient of 1.143 figure 7. calibration experiment at low impact force (approx. 15 kn) with deviation corresponding to force ratio coefficient of 1.048 in recommended installation, shown in figure 4, the load cell is preloaded with beryllium-copper (be-cu) thread stud which is supplied by the manufacturer. this configuration allows a portion of the force to pass through the stud. according to the manufacturer, the be-cu stud may cause deviation of 5 % of the measured force. in our case, steel stud was used which could cause a further increase in deviation. the total deviation is also dependent on the applied preload force. if this force is less than 20 % fso, calibration of the load cell is also necessary. 4. results all measured data are plotted in figure 8 as load cell transducer non-linear output dependency on applied force. because of low level of applied preload force, the strong non-linearity was observed throughout the measured range. as shown in figure 3 under 20 % of fso the load cell has a strong nonlinear characteristics. in that case it is necessary to set the correction coefficient depending on the applied force to obtain the correct value. one possibility is to divide the nonlinear region into sufficiently small sections and linearize them to obtain the corresponding conversion factor for each of them based on the strain gauge signal. another possibility is to define the function approximating the non-linearity of the measured forces for the required measuring area to determine the corresponding value of the conversion coefficient according to this function. the load cell calibration is strongly dependent on the required range of measured forces with respect to the total measuring range of the load cell and the maximum technically achievable preload force. the situation can be divided into two basic cases: 1. preload force itself or with applied load is more than 20 % of fso output has quasi-linear dependency. one correction coefficient is used. the measurement can be strongly influenced by power transmission through the screw. 34 vol. 25/2019 direct measurement of reaction forces figure 8. measured nonlinearity of the force transducer 2. preload force with applied load is less than 20 % of fso output has non-linear dependency. it is necessary to perform calibration over the entire measured range and to determine the discrete or continuous correction coefficients. the effect of the preload bolt is considerably lower. 5. conclusion the performed experiments show that it is possible to use the ohpb for calibration and subsequent measurement with a piezoelectric force transducer. the load cell accuracy is limited by the type of strain gauges used to calibrate it.therefore, it is not possible to reach higher precision with load cell than with strain gauges. however, the reliability of strain gauges and their service life is strongly influenced by the process and quality of the instrumentation, causing a number of failures especially in high strain-rate tests. in contrast, the load cell exhibits high reliability and does not suffer significantly due to cyclic loading. the load cell signal quality is also substantially higher compared to the strain gauges, and in some measurements it can be critical for a reliable evaluation of the experiment. the main objective is to choose a suitable load cell with respect to the measured range of deformation forces. after selecting a suitable working range of the load cell, it is necessary to ensure that the maximum permissible pressure force will not be exceeded and the reflected tensile deformation waves are eliminated to protect the load cell from possible damage. these requirements can be very difficult to achieve with devices operating on the shpb principle. if a suitable load cell is selected, the quasi-linearity of the output values can be achieved even for measurements of small forces. in cases where it is not possible to use suitable load cell due to the price or installation (max. diameter of preload stud), the methods proposed in this work can be used. acknowledgements the research was supported by the czech science foundation (project no. 19-23675s) and the internal grants of the czech technical university in prague (project no. sgs18/153/ohk2/2t/16 and sgs18/154/ohk2/2t/16). all the financial support is gratefully acknowledged. references [1] t. fila, p. zlamal, j. falta, et al. testing of auxetic materials using hopkinson bar and digital image correlation. epj web of conferences 183:02045, 2018. doi:10.1051/epjconf/201818302045. [2] y. fujii. proposal for a step response evaluation method for force transducers. measurement science and technology 14:1741, 2003. doi:10.1088/0957-0233/14/10/301. [3] j. falta, t. fila, m. adorna, p. zlamal. optimization of hopkinson bar instrumentation for cellular and low impedance materials. engieneering mechanics, 25th international conference, svratka, czech republic pp. 97–101, 2019. [4] d. van nuffel, j. peirs, i. de baere, et al. calibration of dynamic piezoelectric force transducers using the hopkinson bar technique. 2012. [5] b. metz. optimizing 3-component force sensor installation for satellite force limited vibration testing. 28th space simulation conference extreme environments: pushing the boundaries 2014. [6] p. piezotronics. technical sheet of dynamic force transducer of type 200c20, 2019. http://www.pcb.com/products.aspx?m=200c20. 35 https://doi.org/10.1051/epjconf/201818302045 https://doi.org/10.1088/0957-0233/14/10/301 http://www.pcb.com/products.aspx?m=200c20 acta polytechnica ctu proceedings 25:21–24, 2019 1 introduction 2 testing setup 3 piezoelectric force transducer 3.1 load cell calibration 4 results 5 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0019 acta polytechnica ctu proceedings 26:19–23, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app cement paste containing micronized recycled concrete influence of hardening accelerators on the modulus of elasticity jakub ďureje∗, zdeněk prošek czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jakub.dureje@fsv.cvut.cz abstract. the article deals with the selection of a suitable hardening accelerator for cement composite material for mass production of masonry blocks. the blocks contain cement and finely ground recycled concrete in ratio 1:1. three different hardening accelerators in different quantity were tested for possibility of fast removing formwork. the dynamic modulus of elasticity was measured by non-destructive resonance method to determine the initial strengths. the modulus of elasticity was measured 7 and 12 hours after sample production to determine the increase in initial strength. subsequently, the modulus of elasticity was measured at 1, 7, 14, 21 and 28 days after production of the test specimens. the resulting moduli of elasticity were compared with reference samples. keywords: finely ground concrete, hardening accelerator, recycled concrete, recycled masonry blocks. 1. introduction old waste concrete from damaged or destroyed structures has the risky properties because of different attributes and inhomogeneity of this waste, which are caused by the presence of other materials from building demolitions (bricks, wood, glass, reinforcement, etc.). the quality of the recycled concrete thus depends on the degree and diligence of the recycling of the recycled concrete. unfortunately, waste concrete sorting is negatively reflected in the price of this raw material. recycled concrete is most often used as a base layer under the roadway. for the common application, a recycle fraction greater than 1 mm is used, the use of finer fractions brings many complications and additional financial costs [1, 2]. the paper deals with the measurement of the dynamic modulus of elasticity of samples from the early hardening phase up to the age of 28 days. the samples contain cement and recycled concrete in weight ratio 1:1, mixing water and various hardening accelerators. recycled concrete containing the finest fraction 0/1mm was used, this material was gently ground by high speed mill. the use of this fine fraction, which today has almost no use, is an advantage of this material, because fine recycled concrete fractions are often stored in landfills as waste material, it follows that using of this material is cost effective and environmental friendly. the finely ground recycled concrete is used in cement composite as a filler with certain binding properties [3]. this cementitious composite will be lightened and will be used to produce thermal insulation perimeter blocks. due to the possibility of mass production of blocks on the production line, a rapid increase in initial strength is needed because of the blocks formwork can be removed and used for the production of other blocks in the shortest possible time. for this reason, different hardening accelerators have been added to the mixtures, and the modulus of elasticity has been measured since the early hardening phase [4]. from previous experiments it has been shown that with increasing percentage of recycled concrete in the mixture there is a decrease in compressive strength. to maintain a mixture satisfactory compressive strength, it is planned to add suitable additives to the cement composite, or to reinforce the blocks by scattered reinforcement [5, 6]. 2. materials and specimens a total of 9 sample sets were made, each set containing 3 test specimens. all test samples contained cement and finely ground recycled concrete in ratio 1:1. the water ratio w/c + r was 0.32 for all samples. the sets differed in the solidification and hardening accelerator used and its amount. one set was a reference without the use of a solidification and hardening accelerator. to produce samples, portland cement cem i 42.5 r from radotín from the českomoravský cement company was used. recycled concrete comes from drainage gutters. a fraction 0/16 mm was used, this material was grinded by the high speed mill. the resulting grain size was under 0.25 mm. the fineness of the ground recycled concrete was similar to the cement, which has a surface area of 361 m2/kg. based on previous works, three types of solidification and hardening accelerators has been selected [7, 8]. betodur gold (b) is a chloride-free liquid additive that accelerates the hardening of mortars and concretes containing portland and mixed cements. ekosal l (e) is a liquid alkali-free solidification and hardening 19 https://doi.org/10.14311/app.2020.26.0019 https://ojs.cvut.cz/ojs/index.php/app jakub ďureje, zdeněk prošek acta polytechnica ctu proceedings designation product based on normal dosing density ph [wt. % of cement] [kg/m3] [-] b non-chloride liquid 1.0 – 3.0 1450 ± 30 5 – 7 e non-alkaline liquid 2.0 – 8.0 1320 ± 30 2 – 4 s polycarboxylates 0.2 – 2.5 1050 4.5 table 1. selected hardening accelerators and its properties. micronized hardening water ratio set / material cement concrete recyclate accelerator w/c + r [g] [g] [ml] [-] ref 750 750 0.00 0.32 b5 750 750 3.75 0.32 b10 750 750 7.50 0.32 b15 750 750 11.25 0.32 e5 750 750 11.25 0.32 e10 750 750 7.50 0.32 e15 750 750 11.25 0.32 s1 750 750 11.25 0.32 s2 750 750 7.50 0.32 s3 750 750 11.25 0.32 table 2. composition of the samples. accelerator that is used primarily for dry-sprayed concrete production. unlike commonly used hardening accelerators, it is not strongly alkaline and does not adversely affect the environment and workers’ health. viscocrete 20 gold (s) is a versatile, highly efficient polycarboxylate-based superplasticizer (pce), particularly suitable to produce concrete with the requirement of a rapid rise in initial strength (table 1). the water ratio was chosen based on previous measurements and hardening accelerator quantities as recommended by the manufacturer. an amount of 5, 10 and 15 l of setting and hardening accelerator per m3 of mixture was tested (table 2). for the production of specimens, 40 × 40 × 160 mm laboratory molds were used. the samples were removed from the mold after 6 hours and stored under water in a laboratory environment at a temperature of 21 ± 2 ◦c and a relative humidity of 50 ± 5% for 28 days. 3. experimental methods for all samples, the dynamic modulus of elasticity was measured at the age of 7 and 12 hours and subsequently at the age of 1, 7, 14, 21 and 28 days. a resonance method was used to measure the dynamic modulus of elasticity of samples. advantage of this method is a speed of measurement and repeatability because it is non-destructive method. the method is based on the calculation of dynamic modulus of elasticity and dynamic shear modulus based on measured fundamental resonance frequencies of longitudinal, transverse (bending) and torsional oscillation. the resonance frequencies were measured in the laboratory using the brüel & kjær measuring system, consisting of the brüel & kjær type 4519-003 response sensor, the brüel & kjær type 8206 impact hammer, the brüel & kjær front-end 3560b-120 and the control computer. the sample was placed on the supports of sufficiently low stiffness so that the supports were placed at the nodal points of the first waveform corresponding to the given oscillation – longitudinal, transverse or rotational (fig. 1). the response sensor (s) was attached on the sample to the appropriate position and the oscillation was excited by the impact hammer (b). both signals, response and excitation were recorded and converted to a frequency domain using a fast fourier transform. subsequently, the pulse labshop software version 14.0.1 has calculated the transmission function. based on the resonance peaks in the transmission function, the fundamental resonance frequency of the given oscillation was determined. from the measured resonance frequency, weight, and sample dimensions, the dynamic modulus of elasticity and shear modulus were determined [9]. 20 vol. 26/2020 cement paste containing micronized recycled concrete figure 1. measurement basic natural frequencies from longitudinal (left), transverse (center), and torsional (right) oscillations [10]. 4. results and discussion the dynamic modulus of elasticity value of each set was calculated by the arithmetic mean of all 3 samples in the set. samples were always measured and weighed. the results were calculated on the basis of measured frequencies of longitudinal oscillations (fig. 2). to verify the correctness of the results, the values of the dynamic modulus of elasticity from the transverse oscillation were calculated too. moduli of elasticity from the transverse oscillation are almost identical with the values of the moduli of elasticity from the longitudinal oscillation, which confirms the correctness of the measurement. from the growth of the dynamic modulus of elasticity edyn follows that the maximum increase in edyn is within the first 7 days. between 7 and 28 days edyn growing only slightly. the rapid increase in the modulus of elasticity in the initial hours is important for the mass production of blocks. the highest increase in the modulus of elasticity after 8 hours was achieved by samples containing accelerator of hardening b (fig. 3). with the increasing amount of hardening accelerator, the modulus of elasticity also increased, but the increase in the modulus of elasticity from 10 to 15 l/m3 (7 % modulus increase) was not as pronounced as from 5 to 10 l/m3 (22 % modulus increase). the increase in the modulus of elasticity of b15 samples versus reference samples was 93 %. for samples containing the e hardening accelerator, there was only a slight increase in the modulus of elasticity compared to the reference samples (13 %). we can conclude from the measured values that the modulus of elasticity increased immediately after sample production, which would correspond to the primary purpose of using this hardening accelerator for sprayed concrete. samples with s accelerator of hardening could not be removed from the mold at 8 hours after production due to low strength. samples were removed from the mold after 24 hours. measurement of dynamic modulus of elasticity for samples with hardening accelerator in amount of 5 l/m3 was successfully at the age of 24 hours, in amount of 10 and 15 l/m3 at the age of 7 days. for set edyn [gpa] ref 1.5 b5 2.2 b10 2.7 b15 2.9 e5 1.3 e10 1.7 e15 1.7 table 3. modulus of elasticity after 8 hours. samples s2 and s3 there were a significant decrease in the modulus of elasticity compared to other specimens due to many cracks. the modulus of elasticity after 28 days was approximately the same for all samples, ranging from 20.5 to 20.7 gpa, except for s2 and s3 specimens. for s2 samples, the modulus of elasticity decreased by 36 % compared to reference samples and s3 samples decreased by 20 %. 5. conclusion the paper deals with the selection of the most suitable solidification hardening accelerator to produce concrete blocks containing micronized recycled concrete. three different hardening accelerators (b, e and s) have been selected based on previous measurements and recommendations from manufacturers. to determine the efficiency of the hardening accelerators, the samples were removed from the molds after 6 hours and the dynamic modulus of elasticity was then measured by the resonance method. based on results, it can be concluded that: • the greatest increase in the dynamic modulus of elasticity of the mixtures occurs in the first 7 days of hardening. most then within 24 h. • as the volume of hardening accelerators increases, the modulus of elasticity also increases, but nonlinearly – a suitable amount of hardening accelerator needs to be determined. 21 jakub ďureje, zdeněk prošek acta polytechnica ctu proceedings figure 2. development of dynamic modulus of elasticity from 0 to 28 days. figure 3. development of dynamic modulus of elasticity from 7 to 12 hours. 22 vol. 26/2020 cement paste containing micronized recycled concrete • hardening accelerators b and e do not negatively affect the final modulus of elasticity even in the highest used amount. • the e accelerator of hardening is likely to increase the modulus of elasticity very soon, but it has not been able to achieve sufficient strength for earlier removal from the mold at the used amount. • hardening accelerator s proved to be unsuitable, samples were removed from the mold after 24 hours and the modulus of elasticity was lower than for reference samples. • hardening accelerator b could be used to increase initial strengths, after 8 hours the samples had almost double modulus of elasticity compared to reference samples. 6. acknowledgements this work was financially supported by the czech technical university in prague the project sgs16/201/ohk1/3t/11, by the ministry of industry and trade the research project fv20503. the authors also thank to the center for nanotechnology in civil engineering at the faculty of civil engineering, czech technical university in prague. references [1] a. barbudo, f. agrela, j. ayuso, et al. statistical analysis of recycled aggregates derived from different sources for sub-base applications. construction and building materials 28(1):129–138, 2012. doi:10.1016/j.conbuildmat.2011.07.035. [2] i. vegas, j. a. ibáñez, a. lisbona, et al. pre-normative research on the use of mixed recycled aggregates in unbound road sections. construction and building materials 25(5):2674–2682, 2011. doi:10.1016/j.conbuildmat.2010.12.018. [3] a. barbudo, f. agrela, j. ayuso, et al. mechanical properties of recycled binder/micro-filler cement-based material. construction and building materials 1054(1):234–237, 2014. doi:10.4028/www.scientific.net/amr.1054.234. [4] m. m. sabai, m. g. d. m. b. cox, r. r. mata, et al. concrete block production from construction and demolition waste in tanzania. construction and building materials 72(1):9–19, 2013. doi:10.1016/j.resconrec.2012.12.003. [5] m. lidmila, j. topič, t. plachý, et al. influence of increasing amount of recycled concrete powder on mechanical properties of cement paste. in iop conference series: materials science and engineering, vol. 236, p. 012094. iop publishing, 2017. doi:10.1088/1757-899x/236/1/012094. [6] j. hrůza. design of masonry elements with use of recycled concrete, 2016. [7] e. anastasiou, k. g. filikas, m. stefanidou. utilization of fine recycled aggregates in concrete with fly ash and steel slag. construction and building materials 50:154– 161, 2014. doi:10.1016/j.conbuildmat.2013.09.037. [8] f. cartuxo, j. de brito, l. evangelista, et al. rheological behaviour of concrete made with fine recycled concrete aggregates–influence of the superplasticizer. construction and building materials 89:36–47, 2015. doi:10.1016/j.conbuildmat.2015.03.119. [9] j. ďureje, z. prošek, j. hrůza, p. tesárek. cement matrix containing micronized recycled concrete and the influence of additives. acta polytechnica ctu proceedings 15:20–24, 2018. doi:10.14311/app.2018.15.0020. [10] čsn 73 1372 nedestruktivní zkoušení betonu–rezonačnní metoda zkoušení betonu, 2012. 23 https://doi.org/10.1016/j.conbuildmat.2011.07.035 https://doi.org/10.1016/j.conbuildmat.2010.12.018 https://doi.org/10.4028/www.scientific.net/amr.1054.234 https://doi.org/10.1016/j.resconrec.2012.12.003 https://doi.org/10.1088/1757-899x/236/1/012094 https://doi.org/10.1016/j.conbuildmat.2013.09.037 https://doi.org/10.1016/j.conbuildmat.2015.03.119 https://doi.org/10.14311/app.2018.15.0020 acta polytechnica ctu proceedings 26:19–23, 2020 1 introduction 2 materials and specimens 3 experimental methods 4 results and discussion 5 conclusion 6 acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0081 acta polytechnica ctu proceedings 15:81–87, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app determination of drying time of the rammed earth walls tereza plaček otcovská∗, barbora mužíková, pavel padevět czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: tereza.otcovska@fsv.cvut.cz abstract. the unburned earth is a building material with long history of use. buildings made of unburned clay are all over the world. but at present, unburned earth is minority building material and its properties are not sufficiently investigated. the rammed earth is one of main kind of unburned earth. this paper is focused on drying rate of the rammed earth of known composition. the significant part of the paper is focused on principle of our own research and the main idea of our research is explained here. the second part of the paper is devoted to drying rate experiment and measured results from this experiment. determination of drying time of universal rammed earth walls is the main result of the experiment. keywords: rammed earth, unburned earth, clay, sand, water coefficient, drying rate, drying time. 1. introduction importance of the unburned earth in construction industry is minor, but there are a few reasons why the research of this building material is important. recently, there has been a growing interest in unburned earth which has resulted in a number of scientific articles on this topic. the growing interest in this historic building material is evident in construction industry too (see fig. 1). i can see in my practice of civil engineer that more and more people choose the unburned earth for build of their houses. the growing interest is the first reason why carry out research in the unburned earth (see fig. 2) [1, 2]. figure 1. sales of unfired bricks by heluz cihlářský průmysl v.o.s. the second reason is a need for reconstruction of historic buildings. there are many earth buildings in the world and in the czech republic that need reconstruction. it is very important to have the knowledges of unburned earth properties for correct reconstruction of this buildings [3–7]. third, an earth building is a good way for sustainfigure 2. building of family house from straw and earth in klášterecká jeseň. able development in the construction industry. the energy needed for production of the unburned clay is minimal and the energy for demolition of the earth construction is small, because earth waste can be completely recycled and earth waste does not need to be stored in a waste dump. a good indicator is primary energy investment (pei). pei is an energy used during live cycle of building materials [8–11]. the primary complication about the properties of the unburned earth is in composition of the earth. the properties depend on composition of the earth and the composition of earth can be very variable. currently, a design of load bearing structures from rammed earth needs a lot of civil engineer’s experiences because norms or methodology does not exist for the design of unburned earth constructions in the czech republic. the reliable methodology is very important for easier design of the earth buildings. that are reasons why is necessary to carry out the research focused on the composition of earth mixture and prop81 http://dx.doi.org/10.14311/app.2018.15.0081 http://ojs.cvut.cz/ojs/index.php/app t. plaček otcovská, b. mužíková, p. padevět acta polytechnica ctu proceedings figure 3. the rammed earth. erties of the final earth structures. the reliable dates from the research should be used for the creation of the methodology [3, 12, 13]. the earth constructions can be produced of the earth mixture without or with inorganic binder. the inorganic binder can improve water resistance of unburned earth. but the inorganic binders are not necessary for the earth constructions. there are old historic buildings and there are researches which proved that unburned earth can be sufficiently resistant against the moisture. the important reason for using unburned earth without the inorganic binders is environment and the sustainable development. the unburned earth with inorganic binder is not possible to recycle easily. this are the reasons why our research is focused on the properties of unburned earth without additives [3, 8, 14–16]. the earth construction can be produced from the earth mixture in several ways. the elementary type of the earth construction is the rammed earth. this paper is focused on the rammed earth. the principle of rammed earth technology is pressing down layers of the earth mixture to a formwork (see fig. 3). the rammed earth walls are usually used like load-bearing structures. that is the reason why the mechanical properties and information about drying rate are very important [3, 13, 17]. the main topic of this paper is the drying rate and drying time. information about the drying rate and drying time are important, because the drying is the main way how the unburned earth gains its final mechanical properties. the rammed earth construction are usually load-bearing types and it is necessary to know when the rammed earth construction is possible to load. 2. the idea of our research as stated in the introduction, the norm or relevant instructions about the earth building designing are not available in the czech republic. the quality research about the unburned earth have to be realized for creating of this instructions. it is necessary to analyse unburned earth properties in dependence on its composition. this is the reason why we investigate mark of clay kind of clay gem montmorillonite agl illite kr illite-kaolinite table 1. three basic kinds of clay. figure 4. the test bodies shape of block. the behaviour and properties of the rammed earth from earth mixtures of different composition. the test earth mixtures are produced in our laboratory from the three elementary components (clay, sand, water). clay fulfils a function of a binder, sand fulfils a function of a filling agent and water serves for an activation of clay bonding properties and for good earth mixture processing. the composition of the earth mixtures differ in kind of clay, amount of clay or amount of mix water. the amount of water is defined by the water coefficient (see equation 1). three elemental kinds of clay were used (see table 1). the summary of the earth mixture which have been made can be seen in table 2. w = mw/mc (1) w water coefficient [-] mw weight of water [g] mc weight of clay [g] determination of dependence between the rammed earth properties and the rammed earth composition is our goal. up to now, we have produced 19 types of the earth mixtures and analysed their properties. we focused particularly on the tensile bending strength, the compressive strength, young’s modulus, moisture’s properties, methylene blue test, shrinkage, creeping and drying rate. we used the test bodies of block shape of size 2/2/10 cm and 4/4/16 cm for majority experiments (see fig. 4). we used the test bodies of prism shape with octahedron base and height of 7 cm for measurement of creeping and shrinkage (see fig. 5). we used the test bodies of block shape of different heights with 82 vol. 15/2018 determination of drying time of the rammed earth walls set clay sand/clay weight ratio water coefficient gemi montmorillonite 80/20 0.370 gemii montmorillonite 75/25 0.370 gemiii montmorillonite 75/25 0.295 agli illite 80/20 0.370 agli illite 75/25 0.370 kri illite-kaolinite 80/20 0.370 krii illite-kaolinite 75/25 0.370 kriii illite-kaolinite 80/20 0.400 kriv illite-kaolinite 85/15 0.295 krv illite-kaolinite 85/15 0.370 krvi illite-kaolinite 85/15 0.400 krvii illite-kaolinite 80/20 0.250 krviii illite-kaolinite 80/20 0.290 krix illite-kaolinite 80/20 0.450 krx illite-kaolinite 75/25 0.295 krxi illite-kaolinite 75/25 0.400 krxii illite-kaolinite 70/30 0.295 krxiii illite-kaolinite 70/30 0.370 krxiv illite-kaolinite 70/30 0.400 table 2. summary of the earth mixtures. figure 5. the test bodies shape of prism with octahedron base. aluminium moulds for measurement of drying rate (see fig 6). 2.1. the production of the test bodies the earth mixture is produced from three elemental components (clay, sand, water). the clay is fine powder form (see fig. 7). size of sand grain is to 4 mm. the grain curve of the sand can be seen in the figure 8. the water was used from a water supply line. the three components were mixed and processed into the earth mixture. the earth mixture was pressed figure 6. the test bodies for drying rate experiment. down to the moulds (see fig. 9). the test bodies for mechanical properties experiments and moisture’s properties experiments were demoulded a few minutes after production and inserted to an air-conditioning chamber. the test bodies for drying rate experiments were left in the moulds and inserted to the air-conditioning chamber. the test bodies were left in the air-conditioning chamber until equilibrium moisture content was reached. relative atmospheric humidity in the air-conditioning chamber was 50 % and temperature was 20oc. 83 t. plaček otcovská, b. mužíková, p. padevět acta polytechnica ctu proceedings figure 7. montmorillonite clay. figure 8. the grain curve of the sand. 3. the measure of the drying rate the test bodies in the figure 6 were used for the measurement of the drying rate. the earth mixture was pressed down to to the moulds of four sizes. unidirectional moisture flow of moisture was provided in the test bodies because the test bodies could dry out only by two opposite surfaces. the size of surface for dry out of the test bodies was square shaped of 53 mm long side. the height of the test bodies (moulds) were 5 cm, 6 cm, 7 cm and 8 cm (see fig. 10). the test bodies in the moulds were inserted to the air-conditioning chamber and regularly weighed. weight decrease of the test bodies was monitored. the moisture content was calculated from the weight of the test bodies by the equation 2. w = (mws − mds)/mds (2) w moisture content [-] mws weight of a wet sample [g] mds weight of a dry sample [g] research was focused on the dependence of the drying rate on the thickness of the drying rammed earth (see fig. 10). drying rate was determined for the test bodies. the points of interest in this paper are amount of moisture 50 % and 15 % from original value. from the values figure 9. the pressing down to the mould. obtained the drying time from experiment start we tried to determine the drying time of a universal wall. 3.1. the measured data the set kriii is the set of the earth mixture for this drying rate experiment (see table 2). number of the test bodies of every height were 3. total of 12 test bodies were used for the experiment. the weighing interval was 15 to 60 minutes in the first 12 hours after production of the test bodies. after first 24 hours the weighing interval was extended. the moisture content was calculated by equation 2. the shape of the moisture content curve in time is shown in fig. 11. a total time for this experiment was approximately 7 day (exactly 167.25 hours). the moisture content after 7 days were 0.6 % to 0.9 % (approximate decrease by 85 % from original value of the moisture content). the main decrease of the moisture content was in the first 24 hours. the moisture content after first 24 hours were 2 % to 2.9 % (approximate decrease by 65 % from original value of the moisture content). the moisture content 50 % from original value was reached after 8 to 16 hours depending on the height of the test bodies (see fig. 12). dependence between the drying rate and the thickness of the rammed earth appears to be linear (see fig. 13). time to reach 50 % from original value of the moisture content of different thickness rammed earth is possible determine by drying coefficient fdry. the drying coefficient is defined by equation 3. fdry = ∑n i=1(ti/hi) n (3) fdry drying coefficient [hour/cm] ti time for drying out of the test bodies of a particular size [hour] hi hight of the test bodies of a particular size [cm] n number of the test bodies heights [-] the drying coefficient was calculated 1.8 hour/cm from the moisture content values obtained. base on 84 vol. 15/2018 determination of drying time of the rammed earth walls figure 10. the diagram of the test bodies in the moulds for drying rate experiments. figure 11. the time course of the moisture content. the values obtained is possible to predict that, for example, a 30 cm thick rammed earth wall will dry out from 50 % in approximately 2.25 days. this prediction was done using equation 4. t50 = ww.fdry (4) t50 time for drying out of a rammed earth wall [hour] ww thickness of a rammed earth wall [cm] fdry drying coefficient [hour/cm] the decrease of the moisture content from 85 % from original value was reached after 70 to 123 hours depending on the height of the test bodies (see fig. 14). dependence between the drying rate and the thickness of the rammed earth appears to be linear (see fig. 15). figure 12. the decrease by 50 % from original value of the moisture content. from the time to reach 85 % of the original value of the moisture content of different thickness rammed 85 t. plaček otcovská, b. mužíková, p. padevět acta polytechnica ctu proceedings figure 13. the decrease by 50 % from original value of the moisture content – linear dependence. figure 14. the decrease by 85 % from original value of the moisture content. earth is possible to determine the drying coefficient. the drying coefficient for 85 % decrease of the moisture content was calculated 14.9 hour/cm from the moisture content values obtained (see equation 3). base on the values obtained it is possible to predict that, for example, a 30 cm thick rammed earth wall will dry out from 85 % approximately 18.6 days. this prediction was done using equation 3. 4. conclusions the rammed earth structures are usually load-bearing types and that is the reason why it is necessary to know how long after their production the rammed earth has its final properties (in particular mechanical properties). determination of drying time of the rammed earth of the known composition (set kriii, see table 2) was the goal of the drying rate experiment described in this paper. the decrease of the moisture content was calculated from the measured values of weights of the test bodies. the drying coefficient was determined for the calculation of the drying time of the different thickness rammed earth walls. the points of interest were 50 % decrease of moisture content from original value and 85 % decrease of moisture content from original value. the prediction of drying time is determined by using the drying coefficient. from the obtained values was predicted that 30 cm figure 15. the decrease by 85 % from original value of the moisture content – linear dependence. thick walls will dry out from 50 % of the original moisture content approximately in 2.25 days. the drying coefficient for this value is 1.8 hour/cm. the same walls will dry out from 85 % of original moisture content approximately 18.6 days. the drying coefficient for this value is 14.9 hour/cm. list of symbols w water coefficient [–] mw weight of water [g] mc weight of clay [g] w moisture content [–] mws weight of a wet sample [g] mds weight of a dry sample [g] fdry drying coefficient [hour/cm] ti time for drying out of the test bodies of a particular size [hour] hi hight of the test bodies of a particular size [cm] n number of the test bodies heights [–] t50 time for drying out of a rammed earth wall [hour] ww thickness of a rammed earth wall [cm] acknowledgements the financial support of this experiment by the czech science foundation (gačr project no. 18-10884s) and faculty of civil engineering, czech technical university in prague (sgs project no. sgs16/201/ohk1/3t/11) is gratefully acknowledged. we would like to express our thanks to lb minerals, s.r.o. company for free supply of material necessary for experimental measurement and we would also like to express thanks to heluz cihlářský průmysl v.o.s. for giving statistical data and 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[17] p. walker. rammed earth: design and construction guidelines, vol. 2010. ihs bre press, watford, 2010. 87 http://dx.doi.org/10.1016/j.buildenv.2008.07.001 acta polytechnica ctu proceedings 15:81–87, 2018 1 introduction 2 the idea of our research 2.1 the production of the test bodies 3 the measure of the drying rate 3.1 the measured data 4 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0126 acta polytechnica ctu proceedings 15:126–130, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app assessment of adhesion between mineral aggregate and bituminous binder using digital image analysis jan trejbala, ∗, zdeněk prošeka, b, tereza valentováa a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b university centre for energy efficient buildings of technical university in prague, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: jan.trejbal@fsv.cvut.cz abstract. this work presents how to assess a rate of adhesion between bituminous binder and mineral aggregate. asphalt mixtures composed from grade bitumen 50/70, reference or modified with adhesion promoters based on amines, and aggregate (brant, zbraslav, skuteč – 8-16 mm) were made and then photographed. three adhesion assessment approaches were applied: (i) standardized adhesion visual assessment, (ii) gray level thresholding, and (iii) entropy-based image segmentation, both evaluated from digital images. it was shown that adhesion between both brant and skuteč and reference binder, expressed as a rate of binder-coated area onto aggregate particles, was equal to ca. 50-70 %, while mixture composed from zbraslav exhibited ca. 70-80 %. if adhesion promoters were used, these areas increased in all three cases up to 80-90 %. it was shown that results obtained using visual and entropy segmentation analysis were very similar, while these differed in comparison with gray-level thresholding. keywords: asphalt mixture, bituminous binder, adhesion promoters, interphase interaction. 1. introduction asphalt mixtures are considered to be one of the most important material in many industries, especially in road structure engineering. after all, over 90 % of roads in europe are constructed from asphalt mixtures, as summarized by s. liu et al. [1]. it is therefore obvious that whatever damages of such roads cause many economical, traffic-safety, and traffic-comfort negative consequences. one of the most common cause of pavement structure failures is so-called “stripping”, i.e., lost of adhesion between bituminous binder film and mineral aggregate particles in dry conditions or in presence of water [2]. this phenomenon is also known as “asphalt moisture damage”. once the adhesion between the two materials fails, water migrates between them and thus disallows their direct contact, as illustrated in figure 1. this is associated with reducing of pavement strength, resulting in damages known as potholes and fatigue cracking. therefore, water is considered to be one of the most enemy of asphalt mixtures. in asphalt mixtures, two types of water sources are known [3]: • internal: water is left in asphalt mixture with inadequately dried aggregate or with hot bitumen at the form of steam. • external: water entering pavements made from asphalt mixtures because of poor drainage or highly wet subgrade. in order to avoid the stripping, it is necessary to dry conditions wet conditions aggregate aggregate bitumen film bitumen film water adhesion lost figure 1. adhesion lost between binder and aggregate in presence of water. ensure a strong interphase interaction between the two materials and thus to prevent to water penetration between them. unfortunately, the adhesion often fails due to physical properties of aggregate (surface smooth texture), occurrence of impurities (clay, dust), or unfavorable bitumen and aggregate chemistry. the last mentioned item seems to be the most problematic issue; binder exhibits weak chemical affinity for aggregate, whereas aggregate is typical for its strong affinity for water, especially if contains high amount of acidic type mineral rocks (high silica contents – quartz-based granites). binder can be therefore replaced by water easily [3]. to avoid the phenomenon described above, the adhesion between the two materials can be enhanced using adhesion promoters – synthetic compounds or those from natural resources, mixed in trace amount to binder. their role is to alter interface properties of both binder and aggregate. their function is based on disruption of intermolecular interactions between 126 http://dx.doi.org/10.14311/app.2018.15.0126 http://ojs.cvut.cz/ojs/index.php/app vol. 15/2018 adhesion assessment between mineral aggregate and bituminous binder water and aggregate surface. consequently, contact energy between binder and aggregate is thus lower [4]. with using of adhesion promoters, an essential question arises: how to quantify their function. across both the technical and laboratory praxis, there are commonly used methods assessing adhesion between binder and aggregate that can be, in general, divided into following two groups: • mechanical testing: asphalt mixtures are made in dry conditions and then their mechanical properties are examined (modulus of elasticity, tensile and pressure strength, adhesion peeling tests, etc.). as the next step, those are exposed to water and then their mechanical properties are tested again. such obtained results are mutually compared [5, 6]. • laboratory interphase interaction testing: interaction between the two materials is tested in laboratory conditions, which cannot reflect many variable parameters like aggregate shape, surface, and so on (contact angle measurement, wilhelmy method, etc.) [7–11]. • stripping rate observation: asphalt mixtures are made similarly to procedure described in the previous item but a rate of stripping is then assessed visually, ie., a rate between binder-coated and stripped areas on aggregate particles is observed (rolling bottle test, detachment in boiling water, etc.) [12, 13]. stripping rate observation-based methods provide the most indicating results in terms of applicability and reliability. they are therefore most frequented across technical praxis, as summarized by grönninger et al. [9]. however, an evaluation of these methods is based on subjective assessment of at least two evaluators, who estimate a degree of bitumen-coated area on aggregate visually. to overcome their subjective and so unconscious conclusions, some researchers have developed semi-automatic techniques. källén et al. [14] and lamperti et al. [15] employed methods based on gray level thresholding (glt), assessing yuv or rgb color space. although these techniques seems to be promising, their usage is limited due to bitumen glossy surfaces. specularly reflected light from these surfaces causes over exposed areas in captured images. these areas merge in with light aggregates, so any gltbased software is not able to recognize them correctly as binder-coated. in order to overcome the issues connected with dark aggregate and over-exposed areas within images, software based on entropy image segmentation was employed. the software is able to assess the adhesion between the two materials even from under-exposed images, free from local over-expositions, but still containing detail drawings in dark parts, if captured properly. these images were used for digital image based analysis: gray level thresholding and entropybased image segmentation. moreover, to compare such obtained results, standardized adhesion visual assessment was also executed. 2. asphalt mixtures 2.1. aggregate three types of mineral aggregate, crushed to the fraction of 8-16 mm, were used with respect to the availability of natural aggregate sources in the czech republic. these three representatives are considered to be the most typical aggregate being used for purposes of asphalt mixture production in the czech republic. moreover, their both volume and surface properties are mutually different; their grain is from very-fine to medium coarse and their surface is from sintered to porous. these differences are suitable for adhesion tests because they cover wide spectrum of all possible aggregate properties. their basic characteristics are summarized in table 1. 2.2. bitumen paving grade bitumen 50/70 was used, standardized by [16]. two alternatives of this bitumen were made – reference with no admixtures and modified containing 0.3 wt. % adhesion promoters based on chemical surfactants comprising amines. 2.3. mixture preparation and imaging following technical standard [13], 12.0±0.3 g of binder was heated to 170±5 °c and then mixed together with 300±3 g of aggregate preheated to 160±5 °c. after prescribed 24 hours lasting technological pause, such prepared asphalt mixtures were poured with 60±3 °c stripping water for 60 minutes. subsequently, these mixtures were subjected to visual assessment according to above cited standard. then, they were photographed using dslr camera canon eos 70d equipped with large-area sensor (aps-c 22.2×14.8 mm), canon zoom lens ef-s 17-85 is usm and polarizing circular filter hoya hd. these pictures were used for semi-automatic adhesion assessment. summarization of all mixtures is for clarity mentioned in table 2. 3. adhesion assessment 3.1. visual analysis the visual analysis, executed according to [13], is based on subjective assessment of at least two evaluators, who determine percentage of binder-coated area on aggregate within asphalt mixture (a reside is considered to be stripped areas). their findings are compared to model cases figured in the cited standard. there are estimated 7 such cases (from a to g and from 100 % to less than 20 % of binder-coated areas) of stripping. it has to be stressed out that their findings can be inaccurate. 3.2. grey level thresholding this assessment is based on recognition of brightness and shadows on digital image of asphalt mixtures. the areas being brighter than the dark bituminous binder are considered to be parts of stripped aggregate 127 j. trejbal, z. prošek, t. valentová acta polytechnica ctu proceedings quarry fraction rock grain surface brant 8-16 mm granodiorite, paleo-porphyrite fine porous zbraslav proterozoic spiliness, metaphrash very fine porous skuteč amphibolic-biotic granodiorite medium-coarse sintered table 1. basic properties of aggregate used. mixture aggregate adhesion promoters [wt. % to binder] br brant 0bm 0.3 zr zbraslav 0zm 0.3 sr skuteč 0sm 0.3 table 2. asphalt mixtures summarization. 0 20 40 60 80 100 su rf ac e of a gg re ga te c oa te d by b in de r [% ] mixture a b c d e g f visual assessment grey level thresholding entropy-based image segmentation br srzr c la ss if ic at io n ac co rd in g to c sn e n 7 3 61 61 70 .0 57 .5 73 .9 80 .0 72 .0 82 .9 49 .7 73 .5 70 .0 figure 2. assessment of adhesion between reference bituminous binder and three types of aggregates. 0 20 40 60 80 100 su rf ac e of a gg re ga te c oa te d by b in de r [% ] mixture a b c d e g f visual assessment c la ss if ic at io n ac co rd in g to c sn e n 7 3 61 61 bm zm sm 90 .0 78 .0 93 .7 90 .0 78 .6 93 .8 90 .0 85 .4 97 .5 grey level thresholding entropy-based image segmentation figure 3. assessment of adhesion between modified bituminous binder and three types of aggregates. 128 vol. 15/2018 adhesion assessment between mineral aggregate and bituminous binder particles. as indicated in introduction (chapter 1), such an approach can be misleading; if aggregate particles are rather dark, the software can define them as binder-coated. on the other side, if bitumen surfaces reflect light specularly, their captured picture is relatively bright so the areas of aggregate assessed by the software are then considered to be stripped. 3.3. entropy-based image segmentation similarly to gray level thresholding, this method is also based on the digital image analysis. its main principle rests on local entropy calculation in order to assess roughness of the texture. although this method is computationally more demanding than that based on brightness/shadows recognition, it is still feasible, and primarily, it produces much more accurate outcomes independent on issues relating with light reflection from binder surfaces. consequently, an input image can be slightly under-exposed in order to eliminate over-exposed areas but only to such an extent to dark areas to be readable for the digital camera. the software is available in [17]. 4. results and discussion 4.1. brant it is clear from figure 2 and figure 3 that the adhesion between reference binder and brant aggregate, classified according to [13], was marked as “d” – coated area up to 70 % (stripped edges, stripped small spots on other areas), while that was increased up to 90 %, classification “b” (aggregate particles coated slightly, edges, corners, and small spots may be stripped) if adhesion promoters were used. when assessed by gray level thresholding based software, the examined coated area reached almost to 60 % and 80 % in the case of using reference and modified binder, respectively. while when the adhesion was assessed using entropy-based image segmentation, the area was over than 70 %, respectively 90 %. 4.2. zbraslav the adhesion between binder and zbraslav aggregate was, in accordance with [13], classified as “c” (80 % – stripped edges and corners, residue well coated) and “b”, when reference and modified binder was used, respectively. in the case of glt analysis, the area was between 72-79 %, regardless to binder type. on the other side, when it was assessed using entropybased image segmentation, the area slightly exceeded 70 %, resp. 80 %. the results are summarized in figure 2 and figure 3. 4.3. skuteč the total of binder-coated areas on skuteč aggregate, consistent with methodology referred to in [13], was equal to “d” and “b” in the case of reference and modified binder used, respectively. when glt analysis was employed, that was approximately 50 %, resp. 85 %, while results from entropy-based image segmentation were equal to almost 75 %, resp. fewer 100 %. these results are also summarized in figure 2 and figure 3. moreover, for clarity, captured images of asphalt mixtures made from skuteč aggregate and reference binder and their preview of consequent assessment using the both semi-automatic methods described earlier are shown in figure 4. 4.4. summarization skuteč and brant proved to be insufficient aggregate in terms of adhesion with reference binder. this was probably caused by their rock-origin; these aggregates, composed namely of acidic rocks, are typical for their fissile surface with high surface energy. on the other side, zbraslav exhibited high affinity for bituminous binder. if adhesion promoters were used, all three aggregates were coated by binder significantly better (in some cases by more than 20 %). after results comparison, it was shown that conventional visual analysis provided very similar results to entropy-based image segmentation; their results differed on average by 2.1 % only. on the other side, grey level thresholding analysis showed to be inaccurate, if compared to other two methods. such obtained results differed by up to 15 %. this was attributed to issues connected with difficulty during asphalt mixture imaging. when the light turned out on the shiny surface of binder, it is then reflected back and causes bright areas in an image. these were next assessed as stripped areas by the software. high occurrence of such areas is captured in figure 4. 5. conclusions three types of aggregate at the fraction of 8-16 mm – brant, skuteč, and zbraslav – were covered by reference and by adhesion promoters modified paving grade bitumen 50/70 for production of asphalt mixture samples to be tested and assessed in terms of adhesion between the two materials. the rate of binder coated areas onto aggregate particles was assessed using three different approaches, concretely by: (i) conventional visual analysis executed according to standardized procedure, (ii) grey level thresholding semi-automatic analysis evaluating a ratio between shadows and brightness in digital image of the mixture, and (iii) entropy-based image segmentation evaluating roughness of image texture. following results were found: • skuteč and brant aggregate showed to be unsatisfactory in intarphase interaction with reference binder. binder-coated area was approximately 50-70 % in both cases, depending on the method employed. if adhesion promoters were used, the adhesion increased by ca. 20 %. • zbraslav exhibited high affinity for reference binder. the coated area was equal to ca. 80 % and 90 % in case od reference and modified binder, respectively. 129 j. trejbal, z. prošek, t. valentová acta polytechnica ctu proceedings figure 4. images of aggregate skuteč covered by reference binder, from left: original photo, gray level thresholding-, and entropy-based image segmentation. • grey level thresholding was intended as the least accurate, if compared to other two methods. this method was not able to recognize bright bitumen areas in captured image, caused by light reflection, as binder-coated. acknowledgements this work was supported by the czech science foundation (research project no. 18-13830s) and by the czech technical university (no. sgs16/201/ohk1/3t/11) and the support is gratefully acknowledged. special thanks to tereza horová for research assistance and to dr. václav nežerka for the software development. references [1] s. liu, a. shukla, t. nandra. technological, environmental and economic aspects of asphalt recycling for road construction. renewable and sustainable energy reviews 75:879–893, 2017. doi:10.1016/j.rser.2016.10.080. [2] b. amit. development of methods to quantify bitumen-aggregate adhesion and loss of adhesion due to water – phd thesis. texas a&m university, 2006. [3] n. ahmad. asphalt mixture moisture sensitivity evaluation using surface energy parameters – phd thesis. the university of nottingham, 2011. [4] c. rossi, b. teltayev, r. angelico. adhesion promoters in bituminous road materials: a review. applied science 7(524), 2017. doi:10.3390/app7050524. [5] s. cui, b. blackman, a. kinloch, a. taylor. durability of asphalt mixtures: effect of aggregate type and adhesion promoters. international journal of adhesion and adhesives 54:100–111, 2014. doi:10.1016/j.ijadhadh.2014.05.009. [6] t. valentová, j. altman, j. valentin. impact of asphalt ageing on the activity of adhesion promoters and the moisture susceptibility. transportation research procedia 14:768–777, 2016. doi:10.1016/j.trpro.2016.05.066. [7] c. rossi, p. caputo, n. baldino, et al. effects of adhesion promoters on the contact angle of bitumenaggregate interface. transportation research procedia 14:768–777, 2016. doi:10.1016/j.ijadhadh.2016.07.013. [8] c. rossi, p. caputo, n. baldino, et al. quantitative evaluation of organosilane-based adhesion promoter effect on bitumen-aggregate bond by contact angle test. international journal of adhesion and adhesives 72:117–122, 2017. doi:10.1016/j.ijadhadh.2016.10.015. [9] j. grönninger, m. wistuba, p. renken. adhesion in bi tumen-aggregate-systems. road materials and pavement design 11(4):881–898, 2010. doi:10.1080/14680629.2010.9690311. [10] a. khan, p. redelius, n. kringos. evaluation of adhesive properties of mineral-bitumen interfaces in cold asphalt mixtures. construction and building materials 125:1005–1021, 2016. doi:10.1016/j.conbuildmat.2016.08.155. [11] j. yi, x. pang, d. feng, et al. studies on surface energy of asphalt and aggregate at different scales and bonding property of asphalt-aggregate system. road materials and pavement design 19(5):1102–1125, 2016. doi:10.1080/14680629.2017.1300597. [12] čsn en 12697-11: bituminous mixtures – test methods for hot mix asphalt part 11: determination of the affinity between aggregate and bitumen, praha: czech office for standards, metrology and testing, 2006. [13] čsn en 76 6161: determination for adhesion of asphaltic binders to aggregate, praha: czech office for standards, metrology and testing, 2000. [14] h. källén, a. heyden, k. åström, p. lindh. measuring and evaluating bitumen coverage of stones using two different digital image analysis method. measurement 84:56–67, 2016. doi:10.1016/j.measurement.2016.02.007. [15] r. lamperti, c. lantieri, c. sangiorgi, et al. semiautomatic evaluation of the degree of bitumen coverage on bitumen-coated aggregates. in conference proceeding of 8th rilem international symposium. rilem bookseries, 2016. doi:10.1007/978-94-017-7342-3_2. [16] čsn en 12591: bitumen and bituminous binders – specifications for paving grade bitumens, praha: czech office for standards, metrology and testing, 2009. [17] v. nežerka. software python phase analyzer based on image segmentation [online], available online: http://mech.fsv.cvut.cz/nezerka/software.html. 130 http://dx.doi.org/10.1016/j.rser.2016.10.080 http://dx.doi.org/10.3390/app7050524 http://dx.doi.org/10.1016/j.ijadhadh.2014.05.009 http://dx.doi.org/10.1016/j.trpro.2016.05.066 http://dx.doi.org/10.1016/j.ijadhadh.2016.07.013 http://dx.doi.org/10.1016/j.ijadhadh.2016.10.015 http://dx.doi.org/10.1080/14680629.2010.9690311 http://dx.doi.org/10.1016/j.conbuildmat.2016.08.155 http://dx.doi.org/10.1080/14680629.2017.1300597 http://dx.doi.org/10.1016/j.measurement.2016.02.007 http://dx.doi.org/10.1007/978-94-017-7342-3_2 acta polytechnica ctu proceedings 15:126–130, 2018 1 introduction 2 asphalt mixtures 2.1 aggregate 2.2 bitumen 2.3 mixture preparation and imaging 3 adhesion assessment 3.1 visual analysis 3.2 grey level thresholding 3.3 entropy-based image segmentation 4 results and discussion 4.1 brant 4.2 zbraslav 4.3 skutec 4.4 summarization 5 conclusions acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0053 acta polytechnica 27(0):53–56, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app hardness of nitrided layers treated by plasma nitriding zdeněk joska∗, zdeněk pokorný, jaromír kadlec, zbyněk studený, emil svoboda university of defence, faculty of military technology, department of mechanical engineering, kounicova 65, 662 10 brno, czech republic ∗ corresponding author: zdenek.joska@unob.cz abstract. stainless steels, particularly the austenitic stainless grades are widely used in many industries due to good corrosion resistance, but very poor mechanical properties as surface hardness and wear resistance limit its possible use. plasma nitriding is one of the few ways to increase the surface hardness of these steels, even though this will affect its corrosion resistance. this paper focuses on the description of the mechanical properties of nitrided layers in the two most widespread austenitic stainless steels aisi 304 and aisi 316l. the microstructure and properties of nitrided layers were evaluated by metallography and microhardness measurement. surface properties of nitrided steels were characterized by martens hardness. the results show that plasma nitriding created very hard nitrided layers with thickness about 40 µm and microhardness about 1300 hv0.05. surface hardness measurements have shown that the maximum values for both steels are about 8.5 gpa, but have different behaviour under higher loads, when the aisi 316l nitrided layer began to crack on the surface and sink. keywords: austenitic stainless steel, microhardness, plasma nitriding. 1. introduction plasma nitriding is one of the few technologies that allow us to increase the surface hardness and abrasion resistance of austenitic steels, and thanks to this surface treatment we can use austenitic stainless steels in other more demanding applications in industries such as chemical, petrochemical food or medical articles. [1–5] the issue of nitrided layers formed on austenitic steels has been intensively solved since the 1990s, when the properties of these layers such as their depth, hardness, abrasion resistance and their influence on the corrosion resistance of steels are investigated. [6–10]. another possible application of such treated surfaces is as a transition layer between a soft backing material such as austenitic steels and a very hard and thin pvd coating [11–14]. the aim of this paper is to describe the mechanical properties of nitrided layers formed on aisi 316l and 304 austenitic stainless steels. 2. experiment the aisi 304 steel samples were 50 x 30 mm and 2 mm thick in the unprocessed state, and the 316l steel had a 30 mm diameter and 7 mm thickness. the surface of the samples was ground and polished to a mirror finish and cleaned in an ultrasonic bath. the values of the chemical composition are given in tab. 1. subsequently, the samples were nitrided in plasma furnace rübig pn 60/60 at condition see tab. 2. sputter cleaning was conducted in an ar–h2 atmosphere prior to plasma nitriding with a pulsed dc glow discharge. all samples were cooled after plasma nitriding in vacuum. depth profiles of plasma nitrided layers were measured by gdoes/qdp method. gdoes measurements were performed in leco sa-2000 equipment. calibration of nitrogen: jk41-1n and nsc4a standards were used. confocal laser microscope lext ols 3000 was used for observation of the cross section morphology of nitrided layers. the microhardness profile was measured on the leco amh 47 automatic microhardness tester. the vickers indenter was selected, the load was 0.049 n, dwell time at load was 12 s. the profile of microhardness was measured at 5 locations on each sample. surface hardness of nitrided samples was measured on the instrument zwick zhu 2.5. the martens hardness of nitrided layers in macroscale was determined. vickers indenter was used, load was 196 n, dwell time at load of 12 s. average value was determined from five measurements. 3. results the nitriding process was carried out under specified conditions; methane was added to the nitriding atmosphere during the nitriding process. gdoes measurements (figs. 1, 2) showed that the plasma nitrided layer had a different chemical composition than the layer formed during standard nitriding. in both steels there were 1 zone in the nitrided layer where the amount of carbon increased. in aisi 304 steel, this zone occurred at a depth of about 10 µm below 53 http://dx.doi.org/10.14311/app.2020.27.0053 http://ojs.cvut.cz/ojs/index.php/app z. joska, z. pokorný, j. kadlec et al. acta polytechnica chemical composition (wt%) c mn cr ni si p s mo aisi 304 0.045 1.78 18.7 8.75 0.48 0.030 0.001 aisi 316 l 0.03 1.50 16.97 10.70 0.31 0.030 0.026 2.18 table 1. chemical composition of stainless steels cleaning nitriding temperature (◦c) 520 550 duration (hours) 0.5 8 h flow h2 (l/min−1) 20 8 flow n2 (l/min−1) 2 32 flow ch4 (l/h) 0 1.5 bias (v) 800 530 pulse length (µs) 100 100 pressure (pa) 80 280 table 2. parameters of plasma nitriding process figure 1. gdoes depth profile of nitrided aisi 304steel the surface, and this layer is very well visible in fig. 3. the second significant zone is below the nitrided layer; this significant zone was formed during the diffusion process as carbon atoms are displaced nitrogen atoms. in aisi 316l steel, a zone of higher carbon content was also found at a depth of 15 µm below the surface, but it is not apparent in the cross-section fig. 4. it can be clearly seen from fig. 4 that the diffusion layer formed has a different character at the bottom, not as sharply delimited as the aisi 304 steel. the total layer thickness of the aisi 304 steel was 45 µm. in aisi 316l steel the nitrided layer reached a thickness of 40 µm. figure 2. gdoes depth profile of nitrided aisi 316l steel figure 3. micrographs of the nitrided layer of the aisi 304 steel the thickness and profile of the nitrided layers were measured by microhardness see fig. 5., 6., thickness of both nitrided layers of aisi 304 and aisi 316l were in good agreement with cross section morphology. the highest values of microhardness were similar but for aisi 316l was higher, reached 1383 hv0.05 and for aisi 304 the highest values of microhardness reached 1200 hv0.05. the highest values of surface hardness were the 54 vol. 27/2020 hardness of nitrided layers treated by plasma nitriding figure 4. micrographs of the nitrided layer of the aisi 316l steel figure 5. microhardness profile of nitrided aisi 304 steel figure 6. microhardness profile of nitrided aisi 316l steel same for both of the steels reached 8.5 gpa. as can be seen from the graphs fig. 7. 8., each nitrided layer behaved differently during loading. with the aisi 304 material, the curve to an indentation depth of about 6 micrometres was stable and then gradually decreased with increasing indentation depth. thus, in the case of aisi 316l, the curve was diametrically different; the layer exhibited stability to an indentation depth of about 15 µm and then ruptured and continued to crack even at an increasing indentation depth. figure 7. force displacement curve of nitrided layers figure 8. hardness vs indentation depth of nitrided layers 4. conclusions the results of the experiment showed that plasma nitriding creates a sufficiently deep nitrided layer on both steels that can be used as a transition layer between a soft substrate and a pvd coating. in aisi 316l steel, the nitrided layer has more suitable properties because it achieves higher hardness and thickness of nitrided layer than aisi 304 steel nitrided layer which has lower microhardness and thickness and in addition, at higher loads, the surface of the nitrided layer is broken. acknowledgements the paper has been prepared thanks to the support of the project the development of technologies, design of firearms, ammunition, instrumentation, engineering of materials and military infrastructure “výzbroj (dzro k201)”. references [1] k. lo, c. shek, j. lai. recent developments in stainless steels. materials science and engineering: r: reports 65(4-6):39–104, 2009. doi:10.1016/j.mser.2009.03.001. [2] t. czerwiec, n. renevier, h. michel. low-temperature plasma-assisted nitriding. surface and coatings technology 131(1-3):267–277, 2000. doi:10.1016/s0257-8972(00)00792-1. 55 https://doi.org/10.1016/j.mser.2009.03.001 https://doi.org/10.1016/s0257-8972(00)00792-1 z. joska, z. pokorný, j. kadlec et al. acta polytechnica [3] j. kadlec, m. dvorak. duplex surface treatment of stainless steel x12crni 18 8. strength of materials 40(1):118–121, 2008. doi:10.1007/s11223-008-0031-y. [4] k. lin, x. li, h. dong, et al. surface modification of 316 stainless steel with platinum for the application of bipolar plates in high performance proton exchange membrane fuel cells. international journal of hydrogen energy 42(4):2338–2348, 2017. doi:10.1016/j.ijhydene.2016.09.220. [5] c. li, t. bell. sliding wear properties of active screen plasma nitrided 316 austenitic stainless steel. wear 256(11-12):1144–1152, 2004. doi:10.1016/j.wear.2003.07.006. [6] j. baranowska. characteristic of the nitride layers on the stainless steel at low temperature. surface and coatings technology 180-181:145–149, 2004. doi:10.1016/j.surfcoat.2003.10.056. [7] j. wang, j. xiong, q. peng, et al. effects of dc plasma nitriding parameters on microstructure and properties of 304l stainless steel. materials characterization 60(3):197– 203, 2009. doi:10.1016/j.matchar.2008.08.011. [8] i. lee. the effect of molybdenum on the characteristics of surface layers of low temperature plasma nitrocarburized austenitic stainless steel. current applied physics 9(3):s257–s261, 2009. doi:10.1016/j.cap.2009.01.030. [9] j. stinville, p. villechaise, c. templier, et al. plasma nitriding of 316l austenitic stainless steel: experimental investigation of fatigue life and surface evolution. surface and coatings technology 204(12-13):1947–1951, 2010. doi:10.1016/j.surfcoat.2009.09.052. [10] z. joska, j. kadlec, v. hruby, et al. characteristics of duplex coating on austenitic stainless steel. key engineering materials 465:255–258, 2011. doi:10.4028/www.scientific.net/kem.465.255. [11] z. joska, m. pospíchal, t. mrázková, j. sukáč. mechanical properties of duplex system: zrn coating on plasma nitrided stainless steel. chemické listy 104:322–325, 2010. [12] e. dalibon, d. heim, c. forsich, s. br uhl. mechanical behavior of nitrided 316l austenitic stainless steel coated with a:c-h-si. procedia materials science 9:163–170, 2015. doi:10.1016/j.mspro.2015.04.021. [13] metallic materials – instrumented indentation test for hardness and material parameters – part 1: test method. čsn en iso 14577-1:2015 . [14] j. stinville, c. tromas, p. villechaise, c. templier. anisotropy changes in hardness and indentation modulus induced by plasma nitriding of 316l polycrystalline stainless steel. scripta materialia 64(1):37–40, 2011. doi:10.1016/j.scriptamat.2010.08.058. 56 https://doi.org/10.1007/s11223-008-0031-y https://doi.org/10.1016/j.ijhydene.2016.09.220 https://doi.org/10.1016/j.wear.2003.07.006 https://doi.org/10.1016/j.surfcoat.2003.10.056 https://doi.org/10.1016/j.matchar.2008.08.011 https://doi.org/10.1016/j.cap.2009.01.030 https://doi.org/10.1016/j.surfcoat.2009.09.052 https://doi.org/10.4028/www.scientific.net/kem.465.255 https://doi.org/10.1016/j.mspro.2015.04.021 https://doi.org/10.1016/j.scriptamat.2010.08.058 acta polytechnica 27(0):53–56, 2020 1 introduction 2 experiment 3 results 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.21.0010 acta polytechnica ctu proceedings 21:10–15, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app microscopical and micromechanical features of high performance concrete containing low levels of fly ash vladimír hrbeka, b, ∗, zdeněk prošeka, roman chylíka, lukáš vráblíka a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague, czech republic b institute of theoretical and applied mechanics as cr, v.v.i., prosecká 76, 190 00 prague 9, czech republic ∗ corresponding author: vladimir.hrbek@fsv.cvut.cz abstract. alkali activated fly ash is a widely used admixture to the concrete. due to the pozzolanic activity, the fly ash can replace up to 60 mass percentage. the impact of fly ash on the concrete depend on the level of it content, i.e. in small levels it improves the macro-mechanical properties and durability as well as chemical resistance. on the other side, fly ash admixture negatively influence the set time and cause slow development of strength. therefore, the effect of small fly ash admixture levels on the micro-structure of high performance concrete (hpc) is discussed in this study. keywords: high performance concrete, ccp, fly ash, grid indentation, spectral deconvolution sem, edx, image analysis, structure, effective modulus. 1. introduction the high performance concrete (hpc) became a commonly used building material in civil over last years due to increased demands on construction costs reduction, quality of building materials, prolonged service life of constructions, etc. compare to ordinary concrete, additional admixtures are added to the mixture, such as fly ash, providing specific features of hardened concrete. fly ash (fa) falls into the category of coal combustion products (ccps) that are produced from either hard or brown coal burning in coal-fired power stations. it is obtained from the gases of furnaces fired with coal or lignite at 1100 to 1400◦c by electrostatic or mechanical precipitation of dust-like particles. fa is defined as fine powder, mainly composed of spherical glassy particles. its chemical composition and pozzolanic and/or latent hydraulic properties depend upon the type of boiler and the type of burned coal. in general, fly ash contains significant quantities of crystalline and amorphous silicon dioxide (sio2), aluminum oxide (al2o3), calcium oxide (cao) and traceable amounts of other elemnets, such as magnesium, chrome, mercury etc. [1, 2] according to statistics of european coal combustion products association (ecoba) from 2016 [3], production of coal combustion products (ccps) in eu 15 countries reached 40.33 million tons and was estimated over 105 million tons in eu 28 countries. wast majority of ccps (63.8%, approx. 25.73 million tons) consists of fa and its utilization in the construction industry is currently around 43% (11.4 million tons). individual ccps production in eu 15 group is depicted in fig. 1a. detailed statistics from eu 28 group are not reliable due to vague characterization of ccps and insufficient data collection. statistics from american coal ash association (acaa) obtained in 2016 [4], show overall ccps production of 107.43 million tons, where only 35.2% (37.82 million tons) is represented by fly ash and 59.9% of it (22.63 million tons) is further used in other industries. it is worth mentioning, that overwhelming majority (75.3%, 17.04 million tons) of fly ash in usa is used in blend cement, concrete and concrete products. fly ash specification and comparison to portland cement, according to american standard (astm c618), are summarized in tab. 2. in czech republic, according to association for the district heating of the czech republic (adh cr) and asvep from 2014, approximately 13 million tons per year of ccps are produced, from which 9.24 million tons (71.1% of estimated production) are classified as fly ash (fig. 1b) . compare to eu 15 countries, czech republic utilize most of ccps, 59%, are used as surface mine backfill and just 11% for cement, concrete and masonry production. even though data of average chemical composition with respect to national standard are not available, specific oxide mass percentage of fly ash obtained from different power plants are presented in tab. 2. national standard csn en 197-1 recognizes fly ash, based on chemical composition, as silica (v type) and calcium fly ash (w type). their average chemical composition, according to [1], is presented in tab. 2. silica fly ash particularly contains silicon dioxide (sio2) and aluminum oxide (al2o3), where content of active sio2 should be over 25% of weight. it also includes more than 10 weight% of active calcium oxide and free form cao should not exceed 1.0 weight%. its use, when free cao (lime) is in range of 1.0 to 2.5 weight %, is limited according to the standard. 10 http://dx.doi.org/10.14311/app.2019.21.0010 http://ojs.cvut.cz/ojs/index.php/app vol. 21/2019 re-cycled concrete (a) ccps production in eu15, 2016. (b) ccps production in czech republic, 2014. figure 1. statistics of ccps production. (note: sda semi dry absorption, fgd flue gas desulphurization, fbc fluidized bed combustion) v type fly ash reacts with calcium hydroxide (ca(oh)2), which is present due to cement hydration, while generating silica hydrates and calcium aluminates. in stage of cement hydration initiation can its particles serve as nucleation centers for hydration products crystallization. calcium fly ash consists of 10 – 15 weight % of active cao and more than 25 weight% of active sio2. compare to silica fly ash, the concentration of alkali and sulfates is higher in w type fly ash and, when calcium hydrate and water is present, it displays pozzolanic and/or hydraulic behavior [5, 6]. fly ash replacement at various levels of the cementitious binder and its impact on hardened composite has been observed on macroscopical level in many studies. higher levels of fly ash addition (30 to 50 weight% replacement of cement) have been in past used mostly in massive structures, like foundations and dams, to control heat generation of the mixture hydration. in recent years, high replacement levels (40 to 60 weight%) used in structural applications produced concrete with increased mechanical properties and durability [7]. dependency of fly ash containment (both low and high levels of replacement of cement) and evolution of concrete strength in time are described in [8, 9]. the improvements of fresh concrete mixture containing fly ash addition, specifically the effect on workability and possible decrease of water-cement ratio, as well as impact on segregation of cement paste and aggregate are discussed in [10–12]. positive impact of the fly ash on shrinkage mitigation depends on two major facotrs, size and size distribution of fly ash particles and water requirements on the mixture. as pointed out in [13, 14], application of 50% replacement of cement by fine fly ash led to the shrinkage reduction by 30%. comparably, when coarse fly ash replacement is used, less reduction of drying shrinkage is observed. other aspect of fly ash addition is its positive effect on permeability and chemical resistance concrete composite. as shown in [15–17], low levels of fly ash replacement decrease the permeability up to 10-15% compare to conventional concrete. however, as in case of more than 50 cement weight% replacement, higher permeability (compare to control specimens) can be observed. on the other hand, significant reduction of chloride permeability (i.e. improvement of chemical resistance), as well as impact on pores size and distribution, with each increment of fly ash replacement volume [18]. 2. materials, sample preparation for purpose of this study, 4 mixtures were selected for testing, referential mixture (ref) and 3 modified mixtures containing 10, 20 and 30 weight% fly ash replacement of cement (fa10, fa20, fa30). ingredients of referential mixture consisted of cementitious binder (cem i 42.5 r mokrá), filler (crushed unwashed basalt aggregate of particle sizes 0 4, 4 8 and 8 16 mm), water and poly-carbonate based plasticizer (stachement, stachema). mixture compositions can be found in tab. 1 component ref fa10 fa20 fa30 cem 42.5 r 800.0 720.0 640.0 560.0 aggreg. 0 4 730.0 730.0 730.0 730.0 aggreg. 4 8 390.0 390.0 390.0 390.0 aggreg. 8 16 320.0 320.0 320.0 320.0 fly ash 0.00 80.0 160.0 240.0 water 210.0 197.4 184.8 172.2 plasticizer 32.0 33.8 29.6 29.5 table 1. compositions of tested mixtures in [kgm−3] specimens (cubes of 100 mm dimension) were stored in water in laboratory conditions (20 22◦c) to ensure adequate cement matrix hydration and to reduce rheologic changes and degradation of tested materials. furthermore, samples were prepared, using grinding and polishing media, for purpose of electron microscopy and indentation. due to mechanical and structural heterogeneity of the composite, preparation method reducing selective abrasivity was adopted. by using this method, both adequate roughness of the specimens surface and compact interface transition zone between aggregate and cement matrix were secured. 11 v. hrbek, z. prošek, r. chylík, l. vráblík acta polytechnica ctu proceedings oxide class f class c v type w type detmarovice hodonin portland cement sio2 50.0 40.0 51.4 39.4 50.4 31.6 23.0 al2o3 26.0 17.0 30.0 30.0 25.5 17.0 4.0 f e2o3 7.0 6.0 4.2 2.2 7.9 6.6 2.0 cao 9.0 24.0 1.3 11.6 4.8 29.4 64.0 m go 2.0 5.0 1.4 1.4 2.8 3.7 2.0 so3 1.0 3.0 0.6 3.8 0.7 7.8 2.0 table 2. oxide analysis in weight percentage [%] of fly ashes an portland cement 3. methods and data processing the testing of described high performance concrete specimens was performed with two independent methods, scanning electron microscopy (sem) and grid nano-indentation. in order to overcome heterogeneous complexity, both methods were implemented on two main levels of the compositematrix without basaltic aggregate and transition zone between aggregate and matrix with limited occurrence of clinker. this prevented duplicity of obtained data from different phases in the material with close mechanical and/or structural features (such as aggregate and clinker etc.). 3.1. scanning electron microscopy and image analysis scanning electron microscope mira ii lmu (tescan corp., brno) equipped with epma was used for purpose of microstructure investigation and phase recognition of hpc samples. chemical composition (i.e. elemental analysis) of each phase was established by energy-dispersive x-ray spectroscopy (edx). this method is precise for determination of weight percentage representation of each chemical element in measured area, nevertheless it is greatly time consuming and covers limited percentage of sample. in general, energy of back-scattered electron is dependent on atomic number z of the element which is relevant to brightness of the phase in bse diagram. this led to direct use of image analyses of back-scattered electron micro-graphs which provided effective description of composite matrix. based on rough estimate of number of phases from low number of edx measurement, it possible to determine proper number of phases and their gray-scale range from histogram of several bse sem graphs. furthermore it is possible numerically quantify percentage representation of each phase and/or convert the diagram onto rgb spectrum image for better imagination of material structure. in case of aggregate-matrix transition zone, twostage image analyses was applied on bse diagrams. the sodium-calcium feldspar forming a structure of basalt aggregate is on gray-scale close to phases present in matrix. differential image analysis than consists of main binding phases (c-s-h gel) separation and identification of other phases in next step of the process based on their internal heterogeneity. 3.2. quasi-static indentation and spectral deconvolution elastic micro-mechanical parameters of individual phases were established usnig grid indentation (ti 700 series, hysitron inc.). principle of indendentation technique is based on dependency of probe propagation with respect to the recorded material response. the force-displacement record is further processed and the mechanical properties are calculated from the unloading part of the record. incorporated errors of the measurement, such as creep and visco-elasticity of measured phases, are avoided by appropriate test setup [19–25]. the implemented indents were displacement driven with maximum depth of 150 nm. for both levels of measurement, load function consisting of loading and unloading over 5 seconds with inserted holding time segment lasting 60 seconds in which maximum depth was held to ensure limited impact of material creep. the indents, again for both levels of investigation, were placed in 21 by 21 grind with mutual separation of 10 µm. due to material level separation and indentation setting, grid indentation criteria of heterogeneous material are met [26–30]. effective modulus of each individual phase present in the composite were derived based on spectral deconvolution of indentation data. the process of spectral deconvolution, unlike statistical deconvolution, takes into account mutual interaction of phases (such as interaction of soft c-s-h gel and tough clinker) and data interference. the deconvolution procedure was implemented individually on both levels of investigated material and and overall effective features were established from merged normalized histograms of indentation data. maximum allowed divergence between structural and mechanical percentage representation of measured data did not exceed 2.5%. 4. results the results of previously described methods and data analysis are stated in this section. mechanical features of individual phases, derived by spectral deconvolution of indentation measurements (fig. 2), are summarized in tab. 3. the structural description of the composite, in means of individual phase percentage representation, was acquired using image analysis of sem bse micro-diagrams (tab. 4). 12 vol. 21/2019 re-cycled concrete sample itz ld csh hd csh portlandite clinker aggregate fa10 16.74 ± 2.42 33.45 ± 5.89 54.54 ± 7.63 84.92 ± 14.07 142.13 ± 13.34 154.86 ± 4.89 fa20 16.48 ± 2.12 34.95 ± 5.36 54.99 ± 6.38 88.63 ± 13.16 140.22 ± 17.42 163.32 ± 7.81 fa30 18.80 ± 4.40 35.62 ± 5.87 57.72 ± 6.40 86.17 ± 7.59 127.57 ± 24.85 161.54 ± 8.28 ref 17.34 ± 2.52 31.08 ± 4.59 45.07 ± 2.61 81.60 ± 18.55 128.36 ± 6.23 162.78 ± 5.87 table 3. indentation modulus er [gpa]) of concrete composites containing fly ash replacement figure 2. histograms of indentation modulus er. sample porosity ld csh hd csh portlandite clinker aggregate fa10 1.19 9.61 15.34 11.77 4.20 57.89 fa20 1.56 12.25 13.96 8.50 2.93 60.78 fa30 2.34 13.48 13.79 8.40 2.81 59.18 ref 1.50 15.82 15.43 6.55 4.18 56.52 table 4. percentage representation of individual phases in composite [%] 5. conclusion from the indentation results of composites on microscopical level and phase representation (based on image analysis of sem bse diagrams), following conclusions can be stated: • lower rates of fly ash replacements decreased interfacial transition zone modulus by 3.6 % for fa10 sample and by 5.2 % in case of fa20 sample while 8.4 % increase can be observed for fa30 specimen. • the indentation modulus of both csh phases (low and high density) are increased in all fly ash replacements. the most influential is 30 % fly ash content, where ld csh modulus is up 14.6 % and hd csh up 28.1 %. • the increase of calcium hydrate (portlandite) micromechanical performance by fly ash additions is up to 8.6 % in case of fa20 sample. • except the lowest fly ash replacement (fa10), fly 13 v. hrbek, z. prošek, r. chylík, l. vráblík acta polytechnica ctu proceedings ash addition increase the microand nanoporosity. • the representation of both csh phases is significantly decreased, while calcium hydrate (portlandite) percentage increased (especially in 10 % fly ash replacement sample). the early stage fly ash containing concrete composite project lower macro-mechanical properties, even though the micro-mechanical features of the matrix are elevated with the fly ash addition. the effect can thus be linked to the structural changes of each composite (mainly to increased porosity and lower csh phases representation). the impact of fly ash addition on the interfacial transition zone is not conclusive, due to inconsistent aggregate sizes in the measurements. further investigation of fly ash replacement on the itz is proposed. list of symbols f measured response/force [µn] hi indentation depth [nm] er indentation modulus [gpa] h hardness [gpa] hc contact depth [nm] p d probability density [–] acknowledgements the research is financially supported by czech technical university in prague – project sgs17/168/ohk1/3t/11 and by the czech science foundation research project 17-19463s. the support is gratefully appreciated. references [1] b. pacewska, g. blonkowski, i. wilinska. investigations of the influence of different fly ashes on cement 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acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0113 acta polytechnica ctu proceedings 22:113–122, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app durability of concrete with binder based on sulfocalcic fly ash rostislav šulc∗, matouš krása, petr formáček czech technical university in prague, faculty of civil engineering, department of construction technology, thákurova 7/2077, 166 29 prague 6, czech republic ∗ corresponding author: rostislav.sulc@fsv.cvut.cz abstract. this paper deals with the durability of concrete with binder based on sulfocalcic fly ash. at the beginning, in the theoretical part, there is a research on the frost resistance of concrete, its resistance to chemical thaw substances and measures to extend the durability of the concrete. there is also a part about sulfocalcic fluid fly ash and concrete. practical part is devoted to individual input materials and experimental procedures. there is also described the methodology of the design of the individual mixtures and the results from the tests and their comparison. keywords: fbc fly ash, durability, chemical thaw substances, compressive strength. 1. introduction we often hear topics about the environment, waste management and recycling in the media. the power industry is one of the leading industries in the production of waste materials. it is 13.7 million tons (2014) of the secondary energy product produced in the czech republic in one year [1]. these are wastes from thermal power plants and heat plants, specifically fly ash, slag, ash and products generated during flue gas desulphurization. attempts are currently being made to utilise these products, especially in the construction industry. fluid ash, which as a by-product is produced by fluidized bed combustion, starts to harden with just mixing with water. the work is based on previous publications dedicated to concrete based on sulfocalcic fly ash [2]. the main point is the resistance of these concrete to chemical thaw substances, where concrete of this type did not meet the requirements of the standards. specifically, it is the preparation of a mixture designed for the construction of a concrete road panels in the categories cb i, cb ii and cb iii [3]. 2. concrete resistance against freezing cycles the topic of frost resistance and resistance of concrete against chemical thaw substances is a very actual. individual stages of environmental aggression are defined. it depends on the conditions in which concrete is located and we can simply define the requirements for concrete [5]. among other things, this standard specifies the following requirements for the properties and composition of the concrete mixture: maximal water coefficient, minimal amount of cement, minimal strength class, minimal air pores content and requirements for used aggregate [4]. 2.1. theory of degradation processes due to frost the mechanism of degradation processes is complicated and ambiguous. there are many theories that differ from each other. the cause is the tension which is caused by the increase in volume of changing water to ice. but this statement is very inaccurate. the pore system of the concrete is very complex and consists of a system of macropores (gel, capillary, air) that are differently saturated with water. in these pores water changes in ice at different temperatures. water does not convert to ice even at -30 to -40 °c in some pores. the gradual formation of ice creates a hydraulic pressure, that forces the water into the capillary system. the degradation of the concrete body occurs if the pressure is greater than the tensile strength of the cement matrix. we also have to take consideration of theories based on thermal expansion of aggregate, cement matrix and ice [5]. there are many theories for the degradation process of frost due to chemical thaw substances. it is referred as salt scaling. it must meet the following characteristics to consider any mechanism as the cause of this type of degradation: (a) degradation consists in the gradual removal of small fragments of sealant. (b) worst condition is usage of a solution of ∼ 3% (independent of the type of solution we use). (c) there is no peeling if there is no layer of solution on the surface. (d) degradation does not occur if the minimum temperature is above -10 °c. the amount of damage increases when the minimum temperature drops below -10 °c. (e) aeration improves durability. 113 http://dx.doi.org/10.14311/app.2019.22.0113 http://ojs.cvut.cz/ojs/index.php/app r. šulc, m. krása, p. formáček acta polytechnica ctu proceedings degree of environmental impact xf1 xf2 xf3 xf4 maximal water coefficient w/c [ ] 0.55 0.55 0.50 0.45 minimal strength class c 30/37 c25/30 c30/37 c30/37 minimal amount of cement [kg/m3] 300 300 320 340 minimal air pores content [%] 4.0 4.0 4.0 table 1. requirements for properties and composition of concrete according to en 206 [4]. (f) the salt concentration in the solution layer on the surface is more important than the salt concentration in the pores. (g) susceptibility to salt peeling is unrelated to the susceptibility to internal frost. (h) surface strength determines the ability to resist the peeling of concrete due to salt [6]. the main cause of salt scaling and the mechanism that meets all these characteristics is called glue spalling. this designation is used originally for the glass decoration method, which explains the mechanism of concrete peeling. this method consists of sandblasting and subsequent application of high temperature epoxy. the formed composite is then cooled. it cracks and creates a tensile stress around the perimeter of cracks because of a larger shrinkage of the epoxide. this results in the peeling off thin glass flaps. it is a composite made of concrete itself and a frozen salt solution on its surface in the case of basic concrete. frozen solution starts to shrink five times more than the concrete if the temperature drops below its melting point. the stress results in the rupture of a certain layer of ice. these cracks penetrate the cement sealant and continue parallel to the composite interface. then some of the concrete is peeled off [7]. 2.2. measures to increase the resistance of concrete against chemical thaw substances the simplest measure is not to expose the concrete to such an environment. but this is usually not possible. there are measures that help to increase and protect the concrete from degradation. aerating additives – we increase concrete resistance to chemical thaw substances by aeration. we must keep a spacing factor of less than 300µm to achieve adequate resistance. aeration reduces the bleeding of the concrete and increases surface strength of concrete. bleeding causes a different density of concrete in its cross section. the greatest value occurs at the bottom. these changes in density correspond to the strength changes and concrete has the smallest strength on the surface. the resulting air bubbles will stick to the cement particles and allow them to move upward. initial frost in the resulting air bubbles causes suction in the pore fluid. this suction results in compression of the skeleton of pores, which reduces the stress caused by thermal expansion to one-third. the mixture should have a pores content with a diameter of less than 300µm greater than 1.5% [6–8]. the water coefficient the water factor has a great effect on the durability of the concrete (strength, permeability), decreasing bleeding and increasing strength while decreasing w/c ratio and achieving greater resistance to salt scaling. concrete with low water coefficient (up to 0.3) do not need to be aerated and its bleeding is very small. it corresponds to concrete with w/c = 0.45 and 20% air [6, 8, 9]. a compressive strength of about 40-45 mpa should be achieved for aerated concrete to be resistant against chemical thaw substances at a conventional water coefficient of 0.4 to 0.45 [6]. aging conditions the longer the concrete treatment in wet conditions, the higher the strength will be. exposing the concrete to high temperatures for a short time immediately after mixing results in a decrease in resistance to chemical thaw substances [6]. the high temperature during aging of concrete has a negative effect on the resistance of concrete against chemical thaw substances as it reduces its strength [6]. it has been proven in concrete that a longer wet treatment period has a positive effect on resistance [10, 11]. subsequent drying at room temperature has no significant effect. the resistance will deteriorate due to the weakening of the concrete surface if the temperature is higher. this weakening can also cause drying at relative humidity less than 30% [6]. sealing mixtures “sealers” only delays the onset of degradation. in some cases, subsequently resulting in sudden and significant drop off generation [6]. when testing three different types of sealants, the resistance test was only delayed by moisture absorption due to the created hydrophobic surface [12]. the moisture content of the treated sample was almost equal to the untreated after one day of immersion and the difference was zero after three days. even with the use of similar silicone treatment preparations [13]. this corresponds to avoiding the degradation during the first 5 to 10 days thaw cycles, but this does not mean that the treatment agent prevents moisture from entering. as mentioned above, the treatment preparations only postpone the onset of degradation but do not affect the resistance to chemical thaw substances [6]. 114 vol. 22/2019 durability of concrete with binder. . . 3. concrete with binder based on sulfocalcic fly ash ashes are made from the combustion of coal in thermal power plants. they are non-combustible inorganic admixtures. the amount of these materials increases with decreasing quality of coal. different types of ash are distinguished depending on the type of combustion chamber, such as high temperature or sulfocalcic fly ash [14, 15]. 3.1. fluid combustion this is a modern method of combustion based on the fluidization effect. it consists of heating inert material (ash or quartz sand, etc.) to the temperature of combustion of fuel and the flow of air to achieve a fluid state of bulk material. liquid properties are used in fluidized fireplaces, hence transferring heat between the liquid and the heating surface. it results in a better heat transfer coefficient than in other types. different types of fuels can be burned. a small amount of limestone is added to the fluidized bed to prevent formation of so3 and to reduce the exhaust of nitrogen oxides. fluid boilers are divided into atmospheric and pressure vessels. atmospheric is further divided into stationary fluid boilers and circulating boilers [15–17]. 3.2. sulfocalcic fly ash the difference is both chemical and mineralogical between sulfocalcic fly ash and classic fly ash. fluid fly ash contains up to 20 wt.% so3, up to 15 wt.% free high reactive cao, and occasionally exhibits greater annealing losses of up to 15 wt.%. when the sulfocalcic fly ash is mixed with water solidification and hardening occur due to the presence of anhydrite and free lime. sulfocalcic fly ash cannot be evaluated according to en 450-1 or en 450-2 due to non-fulfilment of the characteristics on fly ash used in concrete and most technical criteria [18]. 4. concrete with binder based on sulfocalcic fly ash experiments were carried out on concrete in which fly ash was used as a binder [2]. specifically, it was a sulfocalcic fly ash from the tisova power plant. part of the series had a binder composed of sulfocalcic bed and fly ash, referred to as eth10. as a filler was used aggregate composed of three fractions (0-4, 4-8, 8-16). mixtures were not aerated and plasticizing additive glenium ace 40 was used in a ratio of 1.575% of fly ash to achieve better consistency. testing was performed on 100 mm cubes. strength compressive strength increases with longer grinding time. between the non-milled binder and the 45-minute milling was a significant difference of 27.38 mpa, as shown in the table (tab. 2) [2]. resistance against chemical thaw substances – a comparison of two series was made. these series differed only in the water coefficient (b-fpc 57 w/c = 0.4575, b-fpc 48 w/c = 0.4025). with a decreasing water coefficient, the resistance to chemical thaw substances increases, despite they achieved almost the same strength [2]. the values of the resistance of the series using eth10 binders (b-fpc: 57, 58, 59, 64) are recorded in the following table (tab. 3). as can be seen, requirement of standard for waste generation 1000 g/m2 at 75 cycles for cb ii or 100 cycles for cb i does not meet any mixture. among other things, with a longer time of grinding the amount of waste decreases. 5. experimental procedures this part will describe the used materials and their properties and the design methodology and composition of individual samples. 5.1. materials aggregate consisting of three fractions is used is a filler: 0/4, 4/8, 8/16. aggregate consists from sand component originating from dobrin and coarse fractions of crushed aggregate, which are from the zbraslav quarry. the percentage of individual fractions was determined using the kennedy bolomey curve [20]. due to the appreciable absence of the sand fraction in the aggregate, the adjustment of the ratios was made even though the aggregate curve crosses the standard centre line (fig. 1). the resulting ratios are listed in the following table (tab. 4). sulfocalcic fly ash fly ash comes from ledvice coal-fired power station. this is a bed fluid fly ash. the sample of non-grinded fly ash was sent to the xrf laboratory (tab. 5). the order of representation of the individual elements is consistent with the eth10 ash used in the previous work. percentage is different up to 4%. the biggest difference is in the sio2 content, which is 3.81% less in eth 10. additionally, eth 10 has 2.67% more tio2, 2.4% more so3 and 2.26% less cao. the remaining elements do not differ by more than 2% [2]. the ash was ground for 45 minutes in the tumbling mill 0m20f consisting of a rotating drum and grinding charge. in this case, the grinding charge is a mixture of spheres of different diameters (�17 mm 20 kg, �30 mm 20 kg, �50 mm 20 kg) with a total weight of 90 kg. the mill is controlled by a computer, where the rotating speed of the drum (rpm) and the time for grinding and emptying are set on the touch screen [21]. a sample was taken from the non-ground ash ele lo 0 and the ground mixture ele lo iv to determine granulometry (tab. 6, fig. 2, fig. 3). eth 10 fly ash, used in previous experiments, had d50 = 230.6 µm in the non-ground state. a value is less than that of non-grinded ele-lo fly ash, which had d50 = 275.4 µm. as well as for the ground fly ash. for eth 10 – iv value of d50 was 9.3µm and ele lo iv only with d50 = 14.1 µm [2]. ingredients superplasticizing and aerating additives were used in the experiment. glenium ace 40 115 r. šulc, m. krása, p. formáček acta polytechnica ctu proceedings series binder compressive strength28 days [mpa] grinding time [min] water coefficient [ ] b-fpc 57 tisova eth 10 62.9 45 0.458 b-fpc 59 tisova eth 10 35.5 0 0.460 b-fpc 64 tisova eth 10 54.9 15 0.458 table 2. features of the b-fpc series 57, 59, 64 [19]. series sample age[days] chemical thaw substances resistance 25 cycles | 50 cycles | 75 cycles | 100 cycles [g] [g/m2] [g] [g/m2] [g] [g/m2] [g] [g/m2] b-fpc 57 153 50.0 2873.4 102.9 5475.4 b-fpc 58 146 14.2 770.1 29.8 1616.2 42.4 2299.6 81.7 4431.1 b-fpc 59 139 189.5 9036.7 471.4 22479.7 b-fpc 60 139 70.2 2575.4 174.5 6401.8 b-fpc 61 118 88.7 2914.9 202.0 6638.2 b-fpc 62 111 21.6 945.3 41.7 1824.9 70.6 3089.6 81.8 3579.7 b-fpc 63 111 362.0 12387.1 694.4 23829.7 b-fpc 64 104 89.7 3860.6 212.4 9141.4 table 3. results of the chemical thaw substances resistance tests [19]. figure 1. design of aggregate according to kennedy bolomey curve. fraction of aggregate quantity [%] 0-4 42 4-8 18 8-16 40 table 4. aggregate composition. ele lo 0 sio2 al2o3 cao so3 fe2o3 tio2 ∑ 34.16% 27.97% 21.48% 7.89% 3.94% 2.56% 98.00% table 5. xrf chemical composition of fly ash ele lo. r = 50% ele-lo-0 ele-lo-iv dr [µm] 275.374 14.067 table 6. granulometry of the sample and ground sample. 116 vol. 22/2019 durability of concrete with binder. . . figure 2. fly ash granulometry – frequency of particle size. figure 3. fly ash granulometry – passing. superplastifying admixture from basf is composed of polycarboxyether polymers. it accelerates the hydration of cement due to its molecular alignment. the recommended effective dosage is from 0.6% to 1.0% of the weight of the cement. specific dosages are chosen depending on the requirements for processability, type of cement and production technology. glenium ace 40 is applied during mixing, ideally after all ingredients and at least 80% of the mixing water have been added [22]. the mischől lp 78 (now referred to as master air 178) is a synthetic aerating additive. it reduces the surface tension of the mixing water. it achieves good plasticisation and better cohesion of the mixture. the dosage depends on the desired workability, the type of cement, the production technology. the effective dose ranges from 0.05% to 1.5% [23]. 5.2. optimizing aeration additives the amount of aeration additives was optimized in series no. 10. a total of 6 mixtures with different ingredient contents were made. aeration were from 0 to 2.5% and the dosage was increased by 0.5%. the mischől lp 78 was used as an aerating ingredient. the individual mixtures were identical. the dose of aerating ingredient was only thing that changed. the partial composition of each mixture is given in the table below (tab. 7). the series no. 12 was selected as a second mixture for the subsequent chemical thaw substances resistance testing. the aeration was carried out in the same way as in the 10th series. variable is again a dose of aerating additives. there is a lower water factor compared to series no. 10, but it is used 0.5% more plasticizing additive. we can compare the effects of these parameters to the final test values. the composition of individual mixtures in series no. 12 is given in the next table (tab. 8). 6. results and discussion this part is devoted to the results of the different series and their mutual comparison, depending on the different composition and properties. these are the results of concrete slump test (consistency), air content in fresh mixture, compressive strength, and resistance of concrete based on sulfocalcic fly ash against chemical thaw substances. 6.1. consistency this test is based on en 12350-2. the results of slump tests are shown in the following graph (fig. 4). the consistency of series no. 10 is as expected. the slump is growing with the increasing amount of mischől lp 78 additive. the steep slope of the curve at the beginning from 0 to 1% is interesting, especially between 0.5% and 1%, where the 105 mm gap is observed. series no. 12 also show this increasing tendency, but there is an abnormality in mixture with an aerosol content of 1.5%. the slump drops from 190 mm to only 90 mm and then increases again to 240 mm at 2%. this special phenomenon has not been explained. it was originally thought to be due to a dry mixer. the mixture was therefore prepared again, but similar slump values were achieved. a greater difference in consistency is achieved with the gradual dosing of the aerating additive in a series with a greater water coefficient and a lower dose of plasticizer. 6.2. air content in fresh mixture to determine the air content in the fresh concrete mixture, a pressure gauge method was used, the procedure of which is described in the standard en 12350-7. this method is based on a known volume of air and its pressure, which is connected in a sealed container with the volume of air in the concrete mixture, which is unknown [24]. the results from the air content measurements in the fresh concrete mixture are shown in the following graph (fig. 5). 117 r. šulc, m. krása, p. formáček acta polytechnica ctu proceedings sample aggregate fly ash water gleniumace 40 mischől lp 78 air content slump 0-4 4-8 8-16 ele-lo-iv [kg] [kg] [kg] [kg] [l] [kg] [kg] [%] [mm] 10_0 13.749 5.892 13.094 8.4 3.864 0.126 0.000 1.8 30 10_0,5 13.749 5.892 13.094 8.4 3.864 0.126 0.042 2.8 75 10_1 13.749 5.892 13.094 8.4 3.864 0.126 0.084 3.7 180 10_1,5 13.749 5.892 13.094 8.4 3.864 0.126 0.126 3.9 200 10_2 13.749 5.892 13.094 8.4 3.864 0.126 0.168 3.3 210 10_2,5 13.749 5.892 13.094 8.4 3.864 0.126 0.210 3.7 220 table 7. composition of mixtures of the series no. 10. sample aggregate fly ash water gleniumace 40 mischől lp 78 air content slump 0-4 4-8 8-16 ele-lo-iv [kg] [kg] [kg] [kg] [l] [kg] [kg] [%] [mm] 12_0 14.272 6.117 13.593 8.4 3.528 0.168 0.000 2.1 190 12_0,5 14.272 6.117 13.593 8.4 3.528 0.168 0.042 2.9 210 12_1 14.272 6.117 13.593 8.4 3.528 0.168 0.084 4.5 190 12_1,5 14.272 6.117 13.593 8.4 3.528 0.168 0.126 5.6 90 12_2 14.272 6.117 13.593 8.4 3.528 0.168 0.168 5.4 240 12_2,5 14.272 6.117 13.593 8.4 3.528 0.168 0.210 7.6 235 table 8. composition of mixtures of the series no. 12. figure 4. series no. 10, 12 dependence of the amount of aerating additive on slump. figure 5. series no. 10, 12 dependence of the amount of aerating additive on air content. 118 vol. 22/2019 durability of concrete with binder. . . the air content rise with the increasing dose of aerating additive in series no. 10. it increases up to 1.5% and due to the too high flowability of the mixture afterwards the measured air content stop increasing. the air content of the mixture is gradually increasing in the series no. 12 as expected. there was a slight decrease in air content in fresh concrete mixture at the 2% of aerating additive content. the final 2.5% dose reached the highest measured value of all the mixtures produced 7.6% of the air content. 7. compressive strength testing the compressive strength of the cubes (150 × 150 × 150 mm) was based on the test procedure in en 12390-3. weighing samples and measuring their height was prior to testing (in the case of samples that increased their volume, all three dimensions were measured). a slider scale was used to calculate the mass of the concrete samples and to determine the surface area [25]. the compression strength results are shown in the graph below (fig. 6). the compressive strength decreases as the amount of aerating additive increases as expected. the figure shows a decreasing trend of both curves. the strength decreases even though the measured air content with the increasing dose of aerating additives did not rise. the reason could be the inaccuracy of a pressure method that did not measure the real air content of the fresh mixture. this theory of inaccuracy of the pressure gauge method is confirmed in the following figure of the dependence of compressive strength on the air content (fig. 7). the samples 10_2 and 10_2,5 were too fluid in the series no. 10., which could also have an impact on low air content and possible measurement inaccuracy. the series no. 12 shows the expected decline in strength with increasing air content. 7.1. resistance to chemical thaw substances testing of samples for resistance to chemical thaw substances was performed by method a, which is described in čsn 73 1326 [26]. the requirement of standard čsn 73 6123-1 for the resistance of concrete against chemical thaw substances in method a is for group cb i 1000 g/m2 after 100 cycles and group cb ii 1000 g/m2 after 75 cycles. all tested samples met both requirements. the measured values were very low, it is necessary to consider measurement inaccuracies. there is seen the effect of the aeration in the case of series no. 10. the sample without aeration additive has the largest wastes. other mixtures are difficult to compare due to too little waste difference. the series no. 12 shows the positive impact of the lower water coefficient compared to the series no. 10. only two mixtures were measured with waste. there was no waste generation after 125 cycles even in the sample without aerating additive. excess doses of mischől lp 78 greater than recommended (1.5%) have a negative effect on resistance against chemical thaw substances. the compressive strengths of the samples were high enough to exceed the requirements of the c30/37 by standard. the dependence of the amount of waste on compressive strength is not observed (fig. 10). there is no apparent relationship between the air content and the measured waste. high air content of the series no. 12 resulted in greater waste generation. there have been cases where is always one mixture with less or zero waste generation and the other with a higher dose of aerating additives with larger waste generation. both mixtures had almost the same air content. we can speculate about the accuracy of the measurement of the pressure gauge method or about the possible negative effect of the high dose of aerating additive. the dependence of increasing doses and decreasing waste amount cannot be seen in the graph. either no waste or less was measured in samples with fewer aeration additives, than in more aerated mixtures. it is important to find an optimal and effective dose of the additive for the individual mixtures. 8. conclusions the following findings were observed during the optimization. the air content measurements using the pressure gauge method appeared to be inaccurate. in the optimized mix was a requirement for consistency of s4, which is slump from 160 to 240 mm. this requirement was met with optimal doses of aerating additives in the resulting series no. 10 and no. 12. with increasing doses of aerating additives, the slump of the mixtures increased. the high consistency of s5 was reached with high doses. the requirement for air content of 4.0% or 4.5% was met only in the series no. 12. the desired value was reached in mixture with 1% of aerating additive. a maximum of air content was 3.7% in series no. 10. the compressive strength was greater than the requirements for concrete of the group cb iii (c25 / 30) and cb i and ii (c30 / 37). similar results were achieved despite the use of fly ash from another power plant. compared to the previous work, where a 63 mpa strength of concrete based on sulfocalcic fly ash with a grinding time of 45 minutes was achieved [18]. the resistance against chemical thaw substances by the method a test showed that both series with different aeration levels meet the requirements for concrete. a maximum waste generation of 29.4 g/m2 was achieved after 100 cycles in the series no. 10. the positive effect of the lower water content and the negative effect due to the too high dose of the aerating agent on the resulting resistance were shown. the previous sulfocalcic fly ash eth10 has reached a higher levels of waste generation, although with the same granulometry and the chemical representation of the individual elements, a possible 119 r. šulc, m. krása, p. formáček acta polytechnica ctu proceedings figure 6. series no. 10, 12 dependence of the amount of aerating additive on compressive strength. figure 7. series no. 10, 12 dependence of air content on compressive strength. figure 8. series no. 10 results of the chemical thaw substances resistance tests. figure 9. series no. 12 results of the chemical thaw substances resistance tests. sample sample age [days] chemical thaw substances resistance 25 cycles | 50 cycles | 75 cycles | 100 cycles | 125 cycles [g] [g/m2] [g] [g/m2] [g] [g/m2] [g] [g/m2] [g] [g/m2] 10_0_5 34 0.0 0.0 0.3 9.8 0.75 29.4 0.75 29.4 0.91 35.7 10_0,5_5 34 0.0 0.0 0.0 0.0 0.00 0.0 0.00 0.0 0.00 0.0 10_1_5 34 0.0 0.0 0.04 1.6 0.13 5.1 0.13 5.1 0.13 5.1 10_1,5_5 34 0.0 0.0 0.04 1.6 0.12 4.7 0.12 4.7 0.27 10.6 10_2_5 34 0.0 0.0 0.03 1.2 0.11 4.3 0.11 4.3 0.11 4.3 10_2,5_5 34 0.0 0.0 0.01 0.4 0.04 1.6 0.04 1.6 0.31 12.2 table 9. series no. 10 results of the resistance to chemical thaw substances tests. 120 vol. 22/2019 durability of concrete with binder. . . sample sample age [days] chemical thaw substances resistance 25 cycles | 50 cycles | 75 cycles | 100 cycles | 125 cycles [g] [g/m2] [g] [g/m2] [g] [g/m2] [g] [g/m2] [g] [g/m2] 12_0_5 34 0.0 0.0 0.0 0.0 0.00 0.0 0.00 0.0 0.0 0.0 12_0,5_5 34 0.0 0.0 0.0 0.0 0.00 0.0 0.00 0.0 0.0 0.0 12_1_5 34 0.0 0.0 0.0 0.0 0.00 0.0 0.00 0.0 0.0 0.0 12_1,5_5 34 0.0 0.0 0.0 0.0 0.00 0.0 0.00 0.0 0.0 0.0 12_2_5 34 0.0 0.0 0.2 6.3 0.50 18.4 0.50 18.4 0.6 21.6 12_2,5_5 34 0.0 0.0 0.0 0.0 0.00 0.0 0.00 0.0 0.3 11.8 table 10. series no. 12 results of the resistance to chemical thaw substances tests. figure 10. dependence of the amount of waste after 125 cycles on the compressive strength. figure 11. dependence of the amount of waste after 125 cycles on the air content. figure 12. dependence of the amount of waste after 125 cycles on the amount of aerating additive. 121 r. šulc, m. krása, p. formáček acta polytechnica ctu proceedings reason could be the different ash morphology. for this reason, further testing would have to be done. acknowledgements this study was part of the research project th02020163 "development and industrial optimisation of manufacturing process of construction materials from coal ash for transport construction", fv30062 "possibilities of utilization of coal-ash from power stations stored at stock piles", sgs16/198/ohk1/3t/11 "ternary binders based on fly ash from circulating fluidized bed combustion", and we are grateful for cooperation with performed tests provided by the company basf stavební hmoty česká republika s.r.o., zkušební laboratoř betonu praha. references [1] a. bilodeau, v. m. malhotra. aci special publication sp-132, chap. concrete incorporating high volumes of astm class f fly ashes: mechanical properties and resistance to deicing salt scaling and to chloride-ion penetration, p. 319–349. american concrete institute, detroit, 1992. [2] z. možíš. fpc betony na bázi upravených sulfátovápenatých popílků. master’s thesis, fakulta stavební, čvut v praze, praha, 2017. [3] čsn 73 6123-1, stavba vozovek – cementobetonové kryty – část 1: provádění a kontrola shody. standard, úřad pro technickou normalizaci, metrologii a státní zkušebnictví, praha, 2014. [4] en 206, beton – specifikace, vlastnosti, výroba a shoda. standard, úřad pro technickou normalizaci, metrologii a státní zkušebnictví, 2014. [5] j. dohnálek. vliv mrazuvzdornosti betonu na jeho povrchové úpravy. beton technologie konstrukce, sanace 3:44–47, 2012. [6] j. j. valenza, g. w. scherer. a review of salt scaling: i. phenomenology. cement and concrete research 37(7):1007–1021, 2007. doi:10.1016/j.cemconres.2007.03.005. [7] j. j. valenza, g. w. scherer. a review of salt scaling: ii. mechanisms. cement and concrete research 37(7):1022– 1034, 2007. doi:10.1016/j.cemconres.2007.03.003. [8] d. n. richardson. review of variables that influence measured concrete compressive strength. journal of materials in civil engineering 3:95–112, 1991. doi:10.1061/(asce)0899-1561(1991)3:2(95). [9] t. c. powers. the properties of fresh concrete. john wiley and sons, inc., new york, 1968. [10] r. gagne, m. pigeon, p. c. aiticin. aci special publication sp-126, chap. deicer salt scaling resistance of high strength concretes made with different cements, p. 185–199. american concrete institute, detroit, 1992. [11] m. lanlois, d. beaupre, m. pigeon, c. foy. aci special publication sp-114, chap. the influence of curing on the salt scaling resistance of concrete with and without silica fume, p. 971–989. american concrete institute, detroit, 1989. [12] t. sedrah, t. pigeon, m. hazrati. durability of building materials and components, vol. 6, chap. the influence of penetrating sealers on the deicer salt scaling resistance of concrete, pp. 487–496. e & fn spon, london, 1993. [13] p. klieger, w. perenchio. silicone influence on concrete resistance to freeze-thaw and de-icer damage. highway research record 18:33–47, 1963. [14] p. fečko, m. kušnierová, h. raclavská, et al. fly ash. ediční středisko všb-tu ostrava, ostrava, 2005. [15] m. topinková. anorganická pojiva. fakulta metalurgie a metriálového inženýrství, všb-tu ostrava, ostrava, 2013. [16] j. andreovský. spalování paliv – kotle, odborné vzdělávání úředníků pro výkon státní správy ochrany ovzduší v české republice seminář č. 2. http://www.ekomonitor.cz/sites/default/files/ obrazky/seminare/ovzdusi/seminar2/10_dil_5b_ tisk_andreovsky.pdf, 2013. [17] v. polach. parní kotle. http://slideplayer.cz/slide/6816422/, 2011. [18] výzkumný ústav maltovin praha, spol. s.r.o., svaz výrobců cementu čr. popílek a jeho použití do betonu. http://www.vumo.cz/wp-content/uploads/2015/05/ popilek_a_jeho_pouziti_do_betonu.pdf, 2002. [19] en 12350-2, zkoušení čerstvého betonu – část 2: zkouška sednutím. standard, úřad pro technickou normalizaci, metrologii a státní zkušebnictví, praha, 2009. [20] en 206-1/z3, beton – část 1: specifikace, vlastnosti, výroba a shoda. standard, český normalizační institut, praha, 2008. [21] brio hranice s.r.o. omílací buben 0m20f. machine information system. [22] basf stavební hmoty česká republika s.r.o. glenium ace 40, technický list, č. 630, 2007. [23] basf stavební hmoty česká republika s.r.o. master air 178, technický list, 2015. [24] en 12350-7, zkoušení čerstvého betonu – část 7: obsah vzduchu – tlakové metody. standard, úřad pro technickou normalizaci, metrologii a státní zkušebnictví, praha, 2009. [25] en 12390-3, zkoušení ztvrdlého betonu – část 3: pevnost v tlaku zkušebních těles. standard, úřad pro technickou normalizaci, metrologii a státní zkušebnictví, praha, 2009. [26] čsn 73 1326, stanovení odolnosti povrchu cementového betonu proti působení vody a chemických rozmrazovacích látek. standard, vydavatelství úřadu pro normalizaci a měření, praha, 1984. 122 http://dx.doi.org/10.1016/j.cemconres.2007.03.005 http://dx.doi.org/10.1016/j.cemconres.2007.03.003 http://dx.doi.org/10.1061/(asce)0899-1561(1991)3:2(95) http://www.ekomonitor.cz/sites/default/files/obrazky/seminare/ovzdusi/seminar2/10_dil_5b_tisk_andreovsky.pdf http://www.ekomonitor.cz/sites/default/files/obrazky/seminare/ovzdusi/seminar2/10_dil_5b_tisk_andreovsky.pdf http://www.ekomonitor.cz/sites/default/files/obrazky/seminare/ovzdusi/seminar2/10_dil_5b_tisk_andreovsky.pdf http://slideplayer.cz/slide/6816422/ http://www.vumo.cz/wp-content/uploads/2015/05/popilek_a_jeho_pouziti_do_betonu.pdf http://www.vumo.cz/wp-content/uploads/2015/05/popilek_a_jeho_pouziti_do_betonu.pdf acta polytechnica ctu proceedings 22:113–122, 2019 1 introduction 2 concrete resistance against freezing cycles 2.1 theory of degradation processes due to frost 2.2 measures to increase the resistance of concrete against chemical thaw substances 3 concrete with binder based on sulfocalcic fly ash 3.1 fluid combustion 3.2 sulfocalcic fly ash 4 concrete with binder based on sulfocalcic fly ash 5 experimental procedures 5.1 materials 5.2 optimizing aeration additives 6 results and discussion 6.1 consistency 6.2 air content in fresh mixture 7 compressive strength 7.1 resistance to chemical thaw substances 8 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0056 acta polytechnica ctu proceedings 26:56–63, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app localization analysis of an orthotropic multi-surface plasticity model under uniaxial stress claudio pagani∗, milan jirásek, martin horák czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: claudio.pagani@fsv.cvut.cz abstract. numerical simulations of masonry structures are often based on continuum macromodelling approaches that need constitutive laws able to phenomenologically reproduce the behavior of the material. to capture the deformation process up to failure, appropriate softening laws are needed to take into account the contraction of the yield stress domain caused by cracking and crushing. it is well known that softening may lead to localization of inelastic strain. this paper focuses on localization analysis of an orthotropic macro-scale model in the framework of multi-surface plasticity, which describes the in-plane behavior of masonry structures. preliminary results reported in this short paper are limited to uniaxial stress states. analytical localization conditions are first derived for uniaxial stress states with principal axes aligned with the material axes of orthotropy. then, localization analysis is extended to an arbitrary angle between the principal stress axes and the axes of orthotropy. keywords: multi-surface plasticity, localization, weak discontinuity, masonry. 1. introduction masonry structures are characterized by two different failure modes, the in-plane and out-of-plane mechanisms. regarding the study of in-plane mechanisms, different strategies could be adopted for numerical simulations of the behavior of masonry structures. among others, continuum macro-modeling is a frequently used approach, in which masonry structures are treated as homogenized continua and, in general, discretized by the finite element method. this kind of modelling needs the choice of a constitutive model, which could be formulated within the framework of nonlinear elasticity, plasticity, damage mechanics, smeared crack models, or their combinations, and its aim is to phenomenologically reproduce the average behavior of masonry. this paper is focused on a study of an orthotropic model in the framework of multisurface plasticity, with the yield function described as the composition of two different surfaces for the tensile and the compressive behavior [1, 2]. both failure surfaces evolve according to hardening/softening laws that may lead to localization of plastic strain. conditions for the onset of localization depend on many factors, such as the stress state, the destabilizing effects of the non-associated flow rule in plasticity, or, in this particular case, the orthotropy of the elastic stress-strain law. in this paper, the localization analysis of the presented macro-model is carried out under the condition of uniaxial stress in tension and in compression. conditions for the onset of localization are derived and preferential directions of the emerging weak discontinuity are evaluated. 2. model description this paper deals with an advanced model for the in-plane failure mechanisms of masonry structures proposed by lourenço [1, 2], which takes into account the anisotropic structure of masonry and is formulated in the framework of multisurface plasticity. the analysis can be reduced to two dimensions when plane-stress conditions are assumed. in the following description, the cartesian axes x and y are considered aligned with the material axes, which, for regular periodic masonry, are the horizontal and vertical ones (bed and head mortar joints, respectively). in order to avoid pathological mesh sensitivity of the results to the element size, model parameters that affect the stress-strain diagrams in tension and compression are adjusted by means of the equivalent length le. in the original paper, the crack band size was related to the area of the elements: le = αh √ ae (1) where αh is 1 or √ 2 for quadratic and linear elements, respectively. this rule, which has no convincing justification, may induce a large error for elongated elements and even for square elements since the crack band size depends also on the orientation with respect to the element sides. in this paper, the projection method is considered, i.e., the effective crack band size is estimated by projecting the element onto the direction perpendicular to the estimated crack band direction [3]. the model requires in total 16 parameters, which are reported in table 1, and are divided into groups governing orthotropic elasticity, tensile and compres56 https://doi.org/10.14311/app.2020.26.0056 https://ojs.cvut.cz/ojs/index.php/app vol. 26/2020 localization analysis of an orthotropic multi-surface plasticity model sive plastic behavior. the meaning of the symbols is described in the following sections. 2.1. orthotropic elasticity the linear elastic stress-strain relation under planestress conditions can be written as σ = deε (2) where stress and strain components are stored in column matrices using the voigt notation: σ =   σxσy σxy   , ε =   εxεy γxy   (3) the elastic stiffness matrix for an orthotropic material reads de = 1 ζ   ex νyxex 0νxyey ey 0 0 0 ζgxy   (4) where ζ = 1 −νxyνyx is an auxiliary parameter. the elastic behavior is characterized by four parameters: ex, ey, νxy, and gxy. the other in-plane poisson’s ratio is determined from symmetry and is given by νyx = νxyeyy/exx. the plane stress condition (σz = 0) leads to the following expression for the out-of-plane normal strain: εz = − 1 ζ [(νxz + νyzνxy)εx + (νxzνyx + νyz)εy] (5) hence, additional two poisson’s ratios (νxz and νyz) have to be known if one is interested in computing the out-of-plane strain. 2.2. plastic behavior in tension an orthotropic rankine type yield surface is considered for the tensile behavior, with orthotropic softening controlled by a single scalar quantity. the corresponding yield function is defined as ft(σ,κt) = (σx −σt,x(κt)) + (σy −σt,y(κt)) 2 + + √( (σx −σt,x(κt)) − (σy −σt,y(κt)) 2 )2 + ατ2xy (6) degradation of the yield stress is governed by a single scalar internal variable κt. the softening laws σt,i(κt) = ft,i exp ( − left,i gft,i κt ) , i ∈{x,y} (7) describe the dependence of the equivalent tensile yield stresses in directions x and y in the form of exponential softening and can be collected into the column matrix qt =   σt,x(κt)σt,y(κt) 0   (8) parameters ft,i and gft,i, i ∈{x,y}, are the tensile strength and the fracture energy in tension corresponding to the i-axis, respectively. to avoid snap-back at the constitutive level, the value of the tensile strength is reduced if the elements are too large, according to the rule ft,i = √ gft,iei le if le > gft,iei f2t,i (9) parameter α = ft,xft,y τ2u,t (10) controls the shear stress contribution to tensile failure, where τu,t is the shear strength in tension. the flow rule is non-associated and is derived from the plastic potential gt(σ,κt) = (σx −σt,x(κt)) + (σy −σt,y(κt)) 2 + + √( (σx −σt,x(κt)) − (σy −σt,y(κt)) 2 )2 + τ2xy (11) the rate of the softening parameter is considered equal to the rate of the maximum principal plastic strain: κ̇t = ε̇p1 = ε̇px + ε̇py 2 + 1 2 √( ε̇px − ε̇py ) + ( γ̇pxy )2 (12) it is assumed that the hardening regimes in tension and compression are uncoupled, i.e., each of the surfaces is associated with its own hardening variable that grows only if the surface is active. under this assumption, the expression of κ̇t reduces to κ̇t = λ̇t (13) where λt is the plastic multiplier associated with the tensile yield surface. in conclusion, there are five model parameters controlling the tensile behavior: two strengths, ft,x and ft,y, two fracture energies, gft,x and gft,y and the parameter α that governs the coupling of shear with the normal stresses. 2.3. plastic behavior in compression the adopted failure surface in compression corresponds to a hill-type yield function, fc = σ2x σ2c,x(κc) + βσxσy + γτ2xy σc,x(κc)σc,y(κc) + σ2y σ2c,y(κc) −1 (14) where β is a parameter that controls the coupling between the normal stresses. parameter γ controls the shear stress contribution to failure and is given by γ = fc,xfc,y τ2u,c (15) 57 c. pagani, m. jirásek, m. horák acta polytechnica ctu proceedings where fm,x and fm,y are the values of uniaxial compression strength in directions x and y, and τu,c is the shear strength in compression. hardening and softening are controlled by a strainlike internal variable κc, defined by the rate equation κ̇c = 1√ σc,x(κc)σc,y(κc) σt ε̇p (16) which, under the assumption of uncoupled hardening regimes in tension and compression, reduces to κ̇c = λ̇c (17) where λc is the plastic multiplier associated with the compressive yield surface. the values of current compressive yield stress in directions x and y, σc,x and σc,y, are evaluated from κc using the hardening-softening law (figure 1) σc,i =   σi,i + (σp,i −σi,i) √ 2κc κp − κ 2 c κ2p κc ≤ κp σp,i + (σm,i −σp,i) ( κc−κp κm,i−κp )2 κp < κc ≤ κm,i σr,i + (σm,i −σr,i) e 2m(κc−κm) κm−κp κm,i < κc (18) where parameter m = σm,i −σp σm,i −σr,i (19) corresponds to parabolic hardening followed by parabolic and later exponential softening. this law contains strain-like parameters κp, which represents the value of κc at the peak, i.e., at transition from hardening to softening, and κm,x and κm,y, which correspond to the transition from parabolic to exponential softening, and also stress-like parameters σi,x and σi,y, which are the initial values of yield stress in directions x and y, σp,x and σp,y, which are the peak values of yield stress, σm,x and σm,y, which are the values of yield stress at transition from parabolic to exponential softening, and σr,x and σr,y, which are the values of residual yield stress. figure 1. hardening/softening law for the compressive yield stress in absence of specific experimental tests, the characteristic yield stress values can be set to σp,i = fc,i σi,i = 1 3 fc,i (20) σm,i = 12fc,i σr,i = 1 10 fc,i (21) the value of the equivalent plastic strain at peak, κp, is considered to be an additional parameter of the model. the value of κm,i depends on the estimated crack band width le (related to the finite element mesh) and is given by κm,i = 75 67 gfc,i lefc,i + κp (22) where gfc,i is the fracture energy in compression corresponding to the i-axis. in order to avoid snapback at the constitutive level, the compressive strength is reduced if the elements are too large, according to the formula fc,i = √ 75 67 gfc,iei le if κm,i < fc,i ei + κp (23) the values of current compressive yield stress in directions x and y could be cast into the column matrix qc =   σc,x(κc)σc,y(κc) 0   (24) an associated flow rule for compression regime is used, which means that the rate of plastic strain is given by ε̇p = λ̇c ∂fc(σ,κc) ∂σ (25) in summary, six model parameters are used for description of the inelastic behavior in compression: compressive strengths fc,x and fc,y, fracture energies gfc,x and gfc,y, coupling parameter between normal stresses in two directions, β, and the equivalent plastic strain value corresponding to the peak, κp. 3. localization analysis in plasticity suppose that plastic loading takes place with only one yield surface activated, and let i be the index of the active surface ( i ∈{t,c}). in the following derivation, a comma denotes a partial derivative with respect to the variable following the comma and no sum over repeated indices is considered. the basic equations characterizing the flow theory of plasticity include the elastic-plastic split, ε = εe + εp (26) the stress-strain law for the elastic part (2), the yield condition fi(σ,qi) = 0 (27) 58 vol. 26/2020 localization analysis of an orthotropic multi-surface plasticity model the flow rule, ε̇p = λ̇igi,σ (28) the definition of hardening/softening variable in the rate form, κ̇i = λ̇i (29) the dependence of the yield strength on the hardening variable qi = hi(κi) (30) and the loading/unloading conditions fi ≤ 0, λ̇i ≥ 0, λ̇ifi = 0 (31) during plastic flow, the yield function must remain equal to zero, and so the rate of its change is also zero. this consideration leads to the consistency condition, λ̇iḟi = 0 (32) from which the elasto-plastic stiffness matrix can be computed. the resulting formula reads dep = de − de : gi,σ ⊗fi,σ : de fi,σ : de : gi,σ −fi,qi ·hi = = de − 1 h de : gi,σ ⊗fi,σ : de (33) where h = fi,σ : de : gi,σ −fi,qi ·hi (34) and hi = qi,κi (35) the condition of incipient weak discontinuity, as inspired by the early works of hadamard [4] and hill [5] and developed, among others, for plasticity [6, 7] and for damage [8, 9], is given by (n ·dep ·n) ·m = 0 (36) where n is the normal to the discontinuity surface, and qep = n ·dep ·n (37) is the elasto-plastic acoustic tensor. according to (36), this tensor is singular at the onset of localization. the eigenvector m corresponding to the zero eigenvalue is called the polarization vector. the elasto-plastic acoustic tensor can be expressed as qep = n ·dep ·n = = n ·de ·n− 1 h n ·de : gi,σ ⊗fi,σ : de ·n (38) which can be recast in a simpler form qep = qe − 1 h b⊗a (39) where qe = n ·de ·n (40) a = n ·de : fi,σ (41) b = n ·de : gi,σ (42) after some calculations, by imposing singularity of the elastoplastic acoustic tensor, the condition of an incipient weak discontinuity (36) becomes −fi,qi ·hi = a ·q −1 e ·b−fi,σ : de : gi,σ (43) since we are in plane-stress conditions and considering the plane of symmetry perpendicular to the z-axis, one can restrict attention to normals n for which n =   n1n2 0   (44) under the normalizing constraint n21 +n22 = 1. considering major and minor symmetries of the orthotropic elastic stiffness tensor and the plane-stress condition, the quantities entering in (43) can be simplified. namely, the orthotropic linear elastic acoustic tensor is given by qe = [ d1111n 2 1 + d1212n22 (d1122 + d1212)n1n2 (d1122 + d1212)n1n2 d1212n21 + d2222n22 ] (45) where d1111 = exx 1 −νxyνyx (46) d2222 = eyy 1 −νxyνyx (47) d1122 = νyxexx 1 −νxyνyx (48) d1212 = gxy (49) other important quantities are evaluated as fi,σ : de : gi,σ = d1111f11i,σg 11 i,σ+ + d1122(f11i,σg 22 i,σ + f 22 i,σg 11 i,σ)+ + 4d1212f12i,σg 12 i,σ + d2222f 22 i,σg 22 i,σ (50) a = [ (d1111f11i,σ + d1122f 22 i,σ)n1 + 2d1212f 12 i,σn2 2d1212f12i,σn1 + (d1122f 11 i,σ + d2222f 22 i,σ)n2 ] (51) b = [ (d1111g11i,σ + d1122g 22 i,σ)n1 + 2d1212g 12 i,σn2 2d1212g12i,σn1 + (d1122g 11 i,σ + d2222g 22 i,σ)n2 ] (52) where fiji,σ and g ij i,σ are the ij components of the derivative of the yielding function and the plastic potential, respectively, with respect to stress tensor. 4. localization analysis for uniaxial stress let us consider the case of uniaxial stress (figure 2) whose direction is inclined by an angle φ with respect to the x-axis that, together with the y-axis, forms the pair of orthotropy axes. let θ be the angle that the generic normal n to the crack surface forms with the material x-axis, so that it can be expressed as n =   cos θsin θ 0   (53) 59 c. pagani, m. jirásek, m. horák acta polytechnica ctu proceedings the stress tensor matrix in the oxy reference frame corresponding to uniaxial tension of magnitude σ̄ and rotated by a generic angle φ with respect to the x-axis is given by σ(σ̄,φ) = qσ̂qt (54) where σ̂ = [ σ̄ 0 0 0 ] q = [ cos φ −sin φ sin φ cos φ ] (55) the value σ̄ should be the one that triggers localization for the given stress angle φ, so it is the value corresponding to the onset of plastic flow and is determined by the condition fi = 0 where fi is yield function given by (6) or (14) for tension or compression, respectively. subsequently, one can substitute the determined value of σ̄ for a given φ into the right-hand side of (43) and obtain an expression depending only on the angle θ. the localization angle θloc is determined as the one that maximizes the right-hand side of (43) and the corresponding value is the critical hardening modulus (hcrit = h(θloc)). the polarization vector m, the eigenvector of the elastoplastic acoustic tensor corresponding to the zero eigenvalue, characterizes the failure mode, ranging from tensile splitting (mode i) with m = n to shear slip (mode ii) with m perpendicular to n. figure 2. localization (n) and polarization (m) vectors for uniaxial stress σ̄ inclined by φ with respect the material x-axis in the case of uniaxial tension and compression along the x or y-axis, the expressions of the critical hardening moduli can be found in a closed form. first, it is useful to define the quantity ξ, common to all the expressions: ξ(θ) = − ( sin4 θ exx + cos4 θ eyy + + exx − 2gxyνxy exxgxy sin2 θ cos2 θ )−1 (56) t in the isotropic case, the previous expression reduces to ξ = −e (57) where e is the young modulus. subsequently, indicating by subscript t or c for tensile or compressive uniaxial stress and by subscript x or y its direction, the expressions of the critical hardening moduli are given by ht,x(θ) = ξ(θ) sin4 θ (58) ht,y(θ) = ξ(θ) cos4 θ (59) hc,x(θ) = ξ(θ) ( βµ 2 cos2 θ + 1 µ sin2 θ )2 (60) hc,y(θ) = ξ(θ) ( µ cos2 θ + β 2µ sin2 θ )2 (61) where µ = √ fcx/fcy. all the critical hardening moduli have the maximum value equal to zero, so localization starts at the onset of plastic flow for tension (since, under tension, abrupt softening occurs) and at the peak of the compressive yield stress for compression. in the case of uniaxial tension (figures 3-4), normal n is aligned with the stress, in fact θ = φ. in the case of uniaxial compression (figures 5-6), for the chosen set of mechanical parameters, the expressions of the critical hardening moduli exhibit two peaks corresponding to two different possible cracks that are symmetric with respect to the compressive stress axis. figure 3. h(θ) for uniaxial tension in x-direction figure 4. h(θ) for uniaxial tension in y-direction the localization analysis is then extended to the generic case of uniaxial tension or compression inclined by an angle φ with respect to the x-axis, with φ ∈ [0,π/2]. the mechanical parameters used are the same chosen by the author of the model for elementary tests under uniaxial stress [1, 2] and are reported in table 1. 60 vol. 26/2020 localization analysis of an orthotropic multi-surface plasticity model elastic parameters tensile parameters exx eyy νxy gxy ft,x ft,y gt,x gt,y α [gpa] [gpa] [-] [gpa] [mpa] [mpa] [n/m] [n/m] [-] 10.0 5.0 0.20 3.0 1.0 0.5 20 10 1.0 compressive parameters fc,x fc,y gc,x gc,y β γ κp [mpa] [mpa] [n/m] [n/m] [-] [-] [-] 10.0 5.0 5000 2500 -1.0 3.0 5·10−4 table 1. mechanical parameters considered for localization analysis figure 5. h(θ) for uniaxial compression in x-direction figure 6. h(θ) for uniaxial compression in y-direction the critical hardening moduli, except for the particular case of uniaxial tension along the material axes, always exhibit two maxima that correspond to the formation of two potential weak discontinuity surfaces. the solid and dashed lines represented in figures 7 to 12 correspond to the aforementioned two solutions. the dependence of the localization angles on the uniaxial stress angle φ for tension is represented in figure 7, while the difference between the localization angles and the tension angle is shown in figure 8. the variation of the inner product n · m between the normal to crack surface and the polarization vector is represented in figure 9, which shows that for φ = {0,π/2} its value is equal to 1 (tensile splitting), whereas for intermediate values of φ the inner product is lower, which corresponds to a misalignment between n and m. it is worth noting that the two graphs are coincident since the polarization vector m1 corresponding to the normal n1 coincides with the other normal n2 and vice versa. figure 7. dependence of θloc on φ for uniaxial tension figure 8. difference between θloc and φ as a function of φ for uniaxial tension figure 9. dependence of n · m on φ for uniaxial tension in the case of uniaxial compression, the dependence of the localization angles on the stress angle is reported in figure 10). from figure 11 one can see that the differences between the localization angles and the stress angle are almost constant and close to the values π/4 and 3π/4. figure 12 shows that the graphs of 61 c. pagani, m. jirásek, m. horák acta polytechnica ctu proceedings figure 10. dependence of θloc on φ for uniaxial compression figure 11. difference between θloc and φ as a function of φ under uniaxial compression the inner products n ·m corresponding to the two crack surfaces coincide as for the tensile behavior and that the polarization vector is orthogonal to the corresponding normal (shear slip) for φ = 0, whereas for different values of φ the inner product increases up to its maximum value for φ = π/2. 5. conclusions localization analysis under uniaxial tensile and compressive stress has been carried out. analytical results in the particular cases of tension or compression aligned with the material principal axes have been obtained and it has been shown that, for tensile stress, the weak discontinuity surface will form in the direction orthogonal to the stress direction while, in the case of compression, two different discontinuity surfaces, symmetric with respect to the stress direction, could develop. the analysis has been extended by varying the angle that the stress forms with the material axes. it has been shown that for both tension and compression two localization angles can be found and some observations related to the dependence of the localization angles and the polarization vectors on the stress angle have been presented. for all the cases described here, the localization condition is for the first time satisfied at the peak of the stress-strain diagram, i.e., the critical plastic modulus associated with the most favorable direction of a potential discontinuity is zero. for this reason, the results are not affected by the size of finite elements. the adjustment of the softening laws dependent on figure 12. dependence of n · m on φ for uniaxial compression the element size according to (7) or (22) affects only the post-localization stage of response. list of symbols α parameter that controls the shear stress contribution to tensile failure β parameter that couples the normal stresses in the compressive failure surface γ parameter that controls the shear stress contribution to compressive failure le crack bandwidth h hardening modulus hcrit critical hardening modulus κp value of the compressive hardening parameter at peak fi,σ gradient of the i-th yield function with respect to the stress tensor gi,σ gradient of the i-th plastic potential with respect to the stress tensor n normal to the discontinuity surface m polarization vector φ uniaxial stress angle with respect to the x-axis θ angle that the generic normal to the discontinuity surface forms with the x-axis θloc localization angle acknowledgements this research has been financially supported by the european regional development fund through the center of advanced applied sciences at the czech technical university in prague (project no. cz.02.1.01/0.0/0.0/16_019/0000778). references [1] p. lourenço. an orthotropic continuum model for the analysis of masonry structures. tech. rep., university of delft, delft, holland, 1995. [2] p. lourenço. an anisotropic macro-model for masonry plates and shells: implementation and validation. ph.d. thesis, delft university of technology, faculty of civil engineering, 1997. [3] m. jirásek, m. bauer. numerical aspects of the crack band approach. computers and structures 2012. doi:10.1016/j.compstruc.2012.06.006. 62 https://doi.org/10.1016/j.compstruc.2012.06.006 vol. 26/2020 localization analysis of an orthotropic multi-surface plasticity model [4] j. hadamard. leçons sur la propagation des ondes. librairie scientifique a. hermann et fils, paris, 1903. [5] r. hill. a general theory of uniqueness and stability in elastic-plastic solids. journal of the mechanics and physics of solids 6(3):236–249, 1958. [6] n. s. ottosen, k. runesson. properties of discontinuous bifurcation solutions in elasto-plasticity. international journal of solids and structures 27(4):401–421, 1991. [7] j. w. rudnicki, j. rice. conditions for the localization of deformation in pressure-sensitive dilatant materials. journal of the mechanics and physics of solids 23(6):371–394, 1975. [8] m. jirásek. mathematical analysis of strain localization. revue européenne de génie civil 11(7-8):977–991, 2007. [9] m. jirásek, m. horák. localization properties of damage models. in computational modelling of concrete structures, pp. 327–336. crc press, 2010. 63 acta polytechnica ctu proceedings 26:56–63, 2020 1 introduction 2 model description 2.1 orthotropic elasticity 2.2 plastic behavior in tension 2.3 plastic behavior in compression 3 localization analysis in plasticity 4 localization analysis for uniaxial stress 5 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2018.15.0036 acta polytechnica ctu proceedings 15:36–40, 2018 © czech technical university in prague, 2018 available online at http://ojs.cvut.cz/ojs/index.php/app effect of viscous damping in numerical calculation of cable-membrane structures by the method of dynamic relaxation miloš huttner∗, petr fajman czech technical university in prague, faculty of civil engineering, thakurova 7, 166 29 prague 6, czech republic ∗ corresponding author: milos.huttner@fsv.cvut.cz abstract. the aim of this article is to assess the speed of convergence of numerical calculation of cable-membrane structures using dynamic relaxation method in the process of finding critical damping pre-calculated on undamped system. the procedure is tested on four different constructions in six numerical cases. the variety of examples is as large as possible to demonstrate the greatest versatility of the test procedure. the efficiency of the procedure is evaluated based on the number of iterations. keywords: cable-membrane structures, dynamic relaxation, critical damping, viscous damping. 1. introduction the design of cable-membrane structures is realized using numerical models of construction. execution of the numerical analysis is not easy. the shape of the structure, unlike conventional structures, is not known in advance. the initial equilibrium shape of the structure must be found during the phase called the form-finding process. the second phase is the static response process. both processes represent a geometrically non-linear case from a mathematical point of view, but each process has its own specifics. different matrix methods have been developed to address both processes (for example force density method [1], update reference strategy [2], transient stiffness method, etc.). the method of dynamic relaxation represents a very suitable alternative to matrix procedures [3]. in addition, the method can be easily used for both phases of the design of cablemembrane structures. consuming cpu time with more unknowns takes approximately linearly, so the method is very well applicable to models with a large number of degrees of freedom [4]. the dynamic relaxation method builds on the principle that the structure under free oscillation is directed to static equilibrium. the fictitious dynamic analysis is used to find static equilibrium. the artificial parameters (masses, a damping factor, a time interval) influence the speed of convergence. the convergence is the fastest if the damping (a viscous damping is considered) is critical [4]. for fast convergence, it is necessary to find the fundamental oscillation frequency, which is not easy for geometrically non-linear cases. one possible way is to use a pre-calculation with an undamped oscillation mode. the aim of this article is to assess the speed of convergence of numerical calculation of cable-membrane structures in the process of finding critical damping pre-calculated on undamped system. the procedure is tested on four different constructions in six numerical cases. both design processes (form-finding and static response) will be included. 2. methodology critical viscous damping coefficient can be found from [4]: c = 4πmf, (1) where c is viscous damping coefficient, m is mass (or vector of masses), f fundamental frequency of oscillation. the fundamental frequency of oscillation can be obtained pre-calculated on undamped system, with c set to zero. based on the coordinate displacement record, the fundamental oscillation frequency is determined by the formula [4]: f = 1 n × ∆t , (2) where ∆t is time interval, n is the number of iterations required to complete one cycle of oscillations in a fundamental mode. as can be seen from the equation 1, the critical damping factor is a multiple of the mass vector. this makes it easy to compare whether another damping coefficient is converging faster. this article investigates the determination of the damping coefficient based on the oscillation of the structure in the undamped mode (according to equation 2) with one cycle of oscillation. the procedure is a comparison with the otherwise set damping factor for a total of six computational cases (on four different structures). the efficiency of the process is evaluated based on the number of iterations. other parameters are considered as follows: the time interval is selected one second. the mass vector is determined based on the stiffness of each node and 36 http://dx.doi.org/10.14311/app.2018.15.0036 http://ojs.cvut.cz/ojs/index.php/app vol. 15/2018 effect of viscous damping in numerical calculation... is set for each node separately (which is one of the most effective methods, for example, see [5]). 3. testing there are a total of four reference examples, which are used together for a total of six cases. the variety of examples is as large as possible to demonstrate the greatest versatility of the test procedure. structures are modeled in 3d space and include the form-finding (ff) and static response (sr) processes. cables are numerically approximated using bar (ff and sr) and cable (sr) elements, see more [6]. membranes are approximated using isoparametric triangular elements (for ff and sr), see [3] or [7]. 3.1. example 1 the structure is taken from [6]. a scheme of the model structure can be seen in figure 1. the span l = 30.48 m. there have been many experiments on this real orthogonal cable net. thus, numerical results can be compared with real-measured values. cable elements are used to model cables. all cable segments have same cross-sectional areas 1.4645 × 10−4 m2 and young’s modulus 8.2737 × 1010 n/m2. the slack lengths of inner cables are 30.419 m, and the remaining cables are 31.76 m. each of the four internal joints 4, 5, 8 and 9 carries a downward concentrated load 35.56 kn. a small self weigh of 1.46 × 10−3 kn/m for cables is considered. the structure has 12 degrees of freedom. only the static response of the design is solved. figure 1. schema of example 1. perspective view. 3.2. example 2 the structure is taken from lewis [4]. the results can be compared for both processes (form-finding and static response). a structure schema is shown in figure 2. the symmetry of initial configuration about the x and y centre axes is maintained. the coordinates of node 29 are x = 12 m, y = 16 m and z = 3 m. bar elements are used to approximate the cables. the structure has 45 degrees of freedom. for the form-finding process, initial geometry is generated using a pre-tension force 90 kn in x direction, (except for the centre cable where it is 120 kn), a 30 kn in the y direction, plus a vertically point load 4.8 kn at all free nodes. for static response process the cross-sectional areas is assumed to be 350 mm2 in the x direction and 120 mm2 in the y direction. the elastic modulus is 160 kn/mm2. the structure is subjected to the further increase in load 2.0 kn at all free nodes. figure 2. schema of example 2. perspective view. 3.3. example 3 it is a fictitious membrane structure. a scheme of the model structure can be seen in figure 3. the structure is supported throughout the circuit and the dimensions are l = 10 m and h = 4 m. to generate mesh, triangular elements with a maximum edge size of 1.25 meters are used. thus, a model with 231 degrees of freedom is created. the model is tested for both design phases (ff and sr). for the ff process, a constant horizontal pretension σ = 2000 kn/m2 is used (the same for directions x and y). thickness of membrane is 0.001 m. for a static response, a vertical flat load fz = 500 n/m2 is considered. the membrane parameters are: the elastic modulus is 30 n/mm2, and the poisson’s ratio is 0.2. figure 3. schema of example 3. perspective view. 3.4. example 4 it is a fictitious cable-membrane structure that has the parameters of a real construction. a scheme of the model structure can be seen in figure 4. the structure is supported in nodes 1 to 6 and both top rings. the cables are around the perimeter of the structure. the dimensions of structure are l = 8 m and h = 3 m and the radius of the two top rings is 1 m. to generate mesh, triangular elements with a 37 miloš huttner, petr fajman acta polytechnica ctu proceedings maximum edge size of 0.7 meters are used. thus, a model with 1623 degrees of freedom is created. the model is tested only for form-finding process. for the ff process, a constant horizontal pretension σ = 1000 kn/m2 is used (the same for directions x and y). the thickness of membrane is 0.001 m. the horizontal pre-tension force 50 kn is considered in all cables. figure 4. schema of example 4. perspective view. 4. results all cases were tested in the author’s script for solving cable-membrane structures using the dynamic relaxation method created in matlab®. the numerical accuracy of all calculations was controlled by a residual limit force set at 0.0001 kn. 4.1. example 1 the numerical finite element model (in final equilibrium state) for the static response is shown in figure 5. fundamental frequency was estimated at f = 0.0189 hz, that corresponds to 0.237 multiple of the mass vector (α = 0.237). figure 5. example 1 (static response) in a final equilibrium state. the figure 6 shows the number of iterations for damping coefficient values, when c = αm. it has been tested for α in the range from 0.05 to 0.6. it can be seen that the critical damping coefficient (red square) was set very well and that the number of iterations is very minimal. 4.2. example 2 the numerical finite element model (in the initial and final state) for the form-finding process is shown in figure 7. fundamental frequency was estimated at figure 6. comparison of the number of iterations for the damping coefficient determined according to equations 1 and 2 (red square) with different damping coefficients for example 1 (sr). figure 7. example 2 (form-finding) in a final equilibrium state. f = 0.0926 hz, that corresponds to 1.163 multiple of the mass vector (α = 1.163). the figure 8 shows the number of iterations for damping coefficient values. it has been tested for α in the range from 0.8 to 1.6. it can be seen that the critical damping coefficient (red square) was set well and that the number of iterations is close to minimum. figure 8. comparison of the number of iterations for the damping coefficient determined according to equations 1 and 2 (red square) with different damping coefficients for example 2 (ff). fundamental frequency for static response was estimated at f = 0.0088 hz, that corresponds to 0.111 multiple of the mass vector (α = 0.111). the figure 9 shows the number of iterations for damping coefficient values. it has been tested for in the range from 0.04 to 0.22. it can be seen that the critical damping coefficient (red box) was set excellent 38 vol. 15/2018 effect of viscous damping in numerical calculation... and that the number of iterations is minimal. figure 9. comparison of the number of iterations for the damping coefficient determined according to equations 1 and 2 (red square) with different damping coefficients for example 2 (sr). 4.3. example 3 the numerical finite element model (in the initial and final state) for the form-finding process is shown in figure 10. fundamental frequency was estimated at f = 0.0151 hz, that corresponds to 0.190 multiple of the mass vector (α = 0.190). figure 10. example 3 (form-finding) in a final equilibrium state. figure 11. comparison of the number of iterations for the damping coefficient determined according to equations 1 and 2 (red square) with different damping coefficients for example 3 (ff). the figure 11 shows the number of iterations for damping coefficient values. it has been tested for in the range from 0.1 to 1.0. it can be seen that the critical damping coefficient (red square) was not set well and that the number of iterations is not close to minimum. fundamental frequency for static response was estimated at f = 0.0077 hz, that corresponds to 0.096 multiple of the mass vector (α = 0.096). the figure 12 shows the number of iterations for damping coefficient values. it has been tested for in the range from 0.04 to 0.22. it can be seen that the critical damping coefficient (red square) was set excellent and that the number of iterations is minimal. figure 12. comparison of the number of iterations for the damping coefficient determined according to 1 and 2 (red square) with different damping coefficients for example 3 (sr). 4.4. example 4 the numerical finite element model (in the initial and final state) for the form-finding process is shown in figure 13. fundamental frequency was estimated at f = 0.0417 hz, that corresponds to 0.523 multiple of the mass vector (α = 0.523). figure 13. comparison of the number of iterations for the damping coefficient determined according to 1 and 2 (red square) with different damping coefficients for example 4 (ff). the figure 14 shows the number of iterations for damping coefficient values. it has been tested for α in the range from 0.2 to 0.9. it can be seen that the critical damping coefficient (red box was set well and that the number of iterations is close to minimum. 5. conclusions this article investigated the determination of the damping coefficient based on oscillation of the struc39 miloš huttner, petr fajman acta polytechnica ctu proceedings figure 14. comparison of the number of iterations for the damping coefficient determined according to 1 and 2 (red square) with different damping coefficients for example 4 (ff). ture in the undamped mode with one cycle of oscillation. the procedure was a comparison with the otherwise set damping factor for a total of six computational cases (on four different structures). the efficiency of the process was evaluated based on the number of iterations. base on the results presented in previous section, it can be concluded that the proposed method of determining the dynamic relaxation damping coefficient is very suitable for cable-membrane structures. when setting a damping coefficient according to the proposed procedure, the minimum number of iterations (or near to minimum) was achieved in four of the six test cases. excellent results are achieved by the static response analysis procedure with a construction close to the final state. it has been shown that only one oscillation cycle can be used to determine the fundamental frequency. using multiple cycles would probably improve the frequency estimate, but the cpu time would increase. it could be tested in further studies. list of symbols c viscous damping coefficient [t s−1] m mass (or vector of masses) [t] f fundamental frequency of oscillation [hz] ∆t time interval [s] n the number of iterations required to complete one cycle of oscillations in a fundamental mode [–] l the span of the structure [m] h the height of the structure [m] α multiple of the mass vector [–] acknowledgements the results presented in this paper are outputs of the research project “sgs18/037/ohk1/1t/11development and application of advanced algorithms for numerical analysis and modelling in mechanics of structures and materials” supported by the czech technical university in prague. references [1] k. linkwitz, h. j. schek. a new method of analysis of prestressed cable networks and its use on the roofs of the olympic game facilities at munich. iabse, 1972. [2] e. r. bletzinger. a general finite element approach to the form finding of tensile structures by the updated reference strategy. int j space struct (14):131–146, 1999. [3] b. h. v. topping, p. iványi. computer aided design of cable membrane structures. saxe-coburg publications, 2007. [4] w. j. lewis. tension structures: form and behaviour. thomas telford, 2003. [5] m. huttner, j. maca, p. fajman. the efficiency of dynamic relaxation methods in static analysis of cable structures. advances in engineering software (89):28–35, 2015. [6] h. deng, q. f. jiang, a. s. k. kwan. shape finding of incomplete cable-strut assemblies containing slack and prestressed elements. computers and structures (83):21–22, 1767–1779, 2005. [7] m. r. barnes. form and stress engineering of tension structures. structural engineering review (6 (3-4)):175–202, 1994. 40 acta polytechnica ctu proceedings 15:36–40, 2018 1 introduction 2 methodology 3 testing 3.1 example 1 3.2 example 2 3.3 example 3 3.4 example 4 4 results 4.1 example 1 4.2 example 2 4.3 example 3 4.4 example 4 5 conclusions list of symbols acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0096 acta polytechnica 27(0):96–100, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app nanoindentation and accelerated property mapping of portland cement paste jiří němeček1a, ∗, jiří němeček2a, jaroslav lukešb a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b czech technical university in prague, faculty of mechanical engineering, department of mechanics, biomechanics and mechatronics, czech republic ∗ corresponding author: jiri.nemecek.1@fsv.cvut.cz (ph.d. candidate) abstract. this paper presents a new nanoindentation mode: accelerated property mapping (xpm) applied to one year old portland cement paste. the principle of the innovative method is similar to classical quasi-static nanoindentation but the time necessary for performing one indent is significantly decreased from minutes to tenths of seconds. the xpm mode is capable of measuring maps of local mechanical properties (reduced modulus, hardness). the method was compared to classical nanoindentation mode. the main advantages of xpm measurement are higher spatial resolution, lower property fluctuation and lower acquisition time. however, xpm is not able to measure some parameters such as creep. also, high speed loading of xpm brings high strain effect which artificially shifts values of reduced modulus to higher values compared to quasi-static loading cases. keywords: cement paste, high speed nanoindentation, strain rate, xpm. 1. introduction nanoindentation is a widely used technique for micromechanical testing of materials. its principle lies in pressing a tip with known geometry to a material. during the indentation process the loading response is measured. classical nanoindentation performs one indent in the order of minutes which leads to long measurement time, especially if hundreds of indents needs to be performed [1]. in recent years, a new mode of testing called accelerated property mapping (xpm) has been developed by bruker-hysitron company. xpm presents a way to perform hundreds of indents in a few minutes and create maps of reduced modulus and hardness with very high spatial resolution. the testing itself is done by three capacitive plate transducer where the tip is attached to piezo ceramic and loading is then applied electrostatically. classical nanoindentation (ni) measurement consists of several steps: moving tip to the position (a few seconds), piezo settlement (80-200 s), drift measurement (40-80 s) and quasi-static loading (typically a few seconds). during the xpm measurement, the piezo settlement and drift measurement are done only before indentation. so, the process of xpm measurement consists only of performing the indent and changing the position. thus, such a process must be very quick in order to avoid changing the drift rate. elastic modulus is then derived from the unloading curve by oliver and pharr theory [2]. the theory is usually implemented as a standard in nanoindenter’s softwares for, both ni and xpm. however, if the ef1ph. d. candidate at ctu in prague 2associate professor at ctu in prague fect of strain rate is taken into account then it can result in an increase of mechanical properties for some materials tested by nanoindentation, i.e., metals [3, 4]. figure 1. typical sem-bse image of cement paste showing individual phases. our study is focused on cementitious materials where the strain-rate effect of macroscopic mechanical properties e.g. an increase of compressive strength is observable on macro-scale [5]. similar effect can be observed on micro-scale using xpm method which affects the results and has not yet been properly studied. however, using xpm method on cementitious materials would be very advantageous. generally, hundreds of indents are needed for the characterization of highly heterogeneous composites [6, 7] and saving time would be a huge benefit. on the micro-scale, it is possible to characterize individual chemical phases by scanning electron microscopy (sem) and energy dispersion x96 http://dx.doi.org/10.14311/app.2020.27.0096 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 portland cement paste mapping ray spectroscopy (eds). hydrated cement paste is mainly composed of calcium-silica-hydrate (c-s-h gel) existing in two characteristic densities, low (ld c-s-h) and high (hd c-s-h), crystalline calcium hydroxide (ch), residual clinker and porosity [7]. the typical layout of the phases is shown in fig. 1. this study aims to compare both methods on a cementitious composite, highlight their advantages and quantify the strain-rate effect of xpm method with rate dependent visco-plastic constitutive model. 2. materials and methods 2.1. sample preparation a hydrated cement paste from portland cement cem i 42.5r (from mokrá, czech republic) was prepared for testing purposes. cement was mixed with water to cement ratio 0.4, cast into cylindrical molds with height of 32 mm and diameter of 20 mm. the samples were demoulded after three days and stored in the water. 28 days old samples were then cut into 5 mm thick slices on precise cut off machine and dried at 50 °c for one day. the surface of the samples was then grinded with sic papers up to grit size of 4000. finally, samples were put into ethanol into ultrasonic cleaner to remove all loose particles. 2.2. nanoindentation the standard quasi-static nanoindentation test was performed using hysitron ti-700 with berkovich diamond tip. seven grids of 10x10 indents with separation of 12.5 µm were prescribed to cover the heterogeneity of the sample. trapezoidal loading diagram with linear loading lasting 3 seconds up to 2 mn, then by 20 seconds holding period, followed by 3 seconds un-loading were prescribed. xpm measurement was performed on hysitron ti-950. maximum load and the separation between indents were carefully chosen to neglect the influence of individual indents. eight grids with 21x21 indents with separation of 2 µm were prescribed on the sample. the relatively small area of 42x42 µm was covered. thus, each map was prescribed next to the residual clinker to include this less present phase into the results. the testing was force controlled with linear loading lasting 0.1 s up to the maximum force of 1.2 mn, followed by holding and un-loading period both lasting also 0.1 s. 2.3. standard evaluation reduced modulus for both testing methods (grid ni and xpm) was evaluated from load displacement curves using the oliver and pharr method [2]. the method is based on sneddon solution [8] of a rigid indenter action on an elastic half space. the indentation modulus er is defined as er = s √ π 2β √ ac , (1) where s is the elastic unloading stiffness, pmax is the maximum load, β is coefficient which considers different geometry of the indentation tip (β = 1.034 for berkovich tip) and ac is the projected contact area. reduced modulus represents a combination of elastic stiffness of the sample and the tip. modulus of elasticity e, of the sample is given by 1 er = 1 −ν2 e + 1 −ν2i ei , (2) where ei and νi are young’s modulus and poisson’s ratio of the indentation tip (for the diamond tip ei = 1141 gpa and νi = 0.07), ν is the poisson’s ratio of the sample (for the cement paste ν = 0.2 ). during holding period it is possible to evaluate short term creep of the material. creep indentation parameter cit is expressed by equation cit(p,t1,t2) = h2 −h1 h1 × 100, (3) which is defined as the relative change of indentation depths h1 and h2 determined between the start and the end holding period at times t1 and t2, respectively. the parameter is dependent on contact force pmax and the duration of holding period. 3. results and discussion 3.1. standard nanoindentation the results of ni gives young’s modulus and hardness maps (fig. 2a) with relatively low spatial resolution. but it still gives valuable information of mechanical properties of individual phases. the range of e values from 14 to 120 gpa and h from 0.4 to 14 gpa were obtained. the frequency density plot of e was constructed from 700 indents, see fig 3. the values of e corresponding to the main peak lie in the range of 23-38 gpa and are mostly composed with ld c-s-h and hd c-s-h gels, the peak of ch was found around 42 gpa and the rest of the indents with young’s modulus higher than 50 gpa belong to residual clinker. these results correspond well with literature [6, 7]. the total time needed for performing one grid of indents (100 indents) was around 500 minutes. 3.2. accelerated property mapping eight maps of mechanical properties were obtained by xpm method. totally 3548 indents were done with very high spatial resolution of 2 µm. the time needed for acquiring one map was about 5 minutes. an example of measured areas by xpm is shown in fig. 2b. young’s modulus and hardness maps corresponds well with sem-bse image. thus, it suits well for identification of individual phases and their boundaries. also, due to high spatial resolution, xpm is capable of identifying individual clinker minerals, red and orange color in fig. 2b with values of e in range of 115-181 gpa. the moduli of other phases were identified as: ld c-s-h (e=30-32 gpa), hd c-s-h (e=40-46 gpa), ch (e=50-56 gpa). these 97 j. němeček, j. němeček, j. lukeš acta polytechnica young’s modulus, e [gpa] 0 20 40 60 80 hardness, h [gpa] a) 90 120 1816030 sem-bse b) 42 μm 112.5 μm 0 1 3 5 7 7 10 1541 figure 2. sem-bse image, young’s modulus, hardness maps measured by a) ni b) xpm. values are higher than the literature results [6, 7]. however, the results correspond well with xpm study of sebastiani et al. [9]. again, the frequency plot of e was constructed, see fig 3. 0 20 40 60 80 100 120 140 160 180 0.00 0.01 0.02 0.03 0.04 0.05 0.06 0.07 shift young’s modulus, e [gpa] fr eq ue nc y de ns ity [-] xpm ni figure 3. frequency plots of ni and xpm with highlighted shift of the main peak. 3.3. indentation creep parameter cit was evaluated from 20 second holding period of ni technique. the results of individual phases are shown in table 1 and a different parameter was found for each individual phase. the highest value was found for c-s-h gel around 9-12%. also, some deformation was found for ch and residual clinker, where it is well known that these phases do not creep. however, during indentation hard particles of clinker are pressed into surrounding matrix of hydrates, which affects the results. again, cit parameter was evaluated from the holding period of xpm. a very interesting fact was found though: the value of cit was 6.8 ± 0.3% xpm ni phase 1.2mn, 2.0mn, 2.0mn, 0.1s,0.2s 3.0s, 3.1s 3.0s, 23.0s ld c-s-h 6.8±0.3% 1.2±0.3% 11.3±3.1% hd c-s-h 6.8±0.3% 1.1±0.3% 9.5±3.3% ch 6.8±0.3% 0.8±0.2% 6.4±3.7% clinker 5.9±0.6% 0.7±0.2% 2.1±1.9% table 1. creep indentation parameter cit calculated in different phases for xpm and ni. regardless of the indentation phase. however, different indentation forces were used. thus, the values are not directly comparable for xpm and ni. the cit was evaluated for 0.1 s holding period from ni. smaller values with different values were found, even if higher force was used. such findings suggest that during xpm, creep of the material is not correctly measured and the results might be affected by creep of the piezo scanner. 3.4. mutual comparison grid ni brings very low spatial resolution with very good absolute values of mechanical properties of ld, hd c-s-h gel and ch. the cit parameter provides comparison of creep deformation between individual phases. xpm gives higher spatial resolution somewhere close to sem-eds maps. the area covered by individual maps is lower than that for ni. however, time needed for obtaining one map is significantly decreased from hours to minutes. the creep parameter 98 vol. 27/2020 portland cement paste mapping 1 μm 1 μm 0.15 0.13 0.11 0.09 0.07 0.05 0.03 0.02 0 unit: (μm/μm) 0.15422 0.15 0.13 0.11 0.09 0.07 0.05 0.03 0.02 0 unit: (μm/μm) 0.15422 b)a) figure 4. equivalent von mises strain under the tip at the end of loading a) at time = 3 s, speed = 0.67mn/s (ni) b) at time = 0.1 s, speed = 20mn/s (xpm). of material is not obtained from the method. quantitatively, xpm provides data comparable to ni but slightly shifted towards higher values. the shift of 14.8% was calculated as the ratio of young’s modulus of the points with the highest frequency density in the histogram measured by ni and xpm, respectively (fig. 3). such a deviation could be explained by strain-rate effect and was further investigated with numerical modeling. 4. numerical modeling the strain rate effect was studied by finite element method (fem) in ansys software as 2d asymmetrical task with transient dynamic solver. the real shape diamond indenter (elastic isotropic:ei = 1141 gpa and νi = 0.07) was brought to contact with an elastoplastic domain with size of 15x15 µm of the cement sample. loading rate effects were captured by viscoplastic perzyna model [10]. the model calculates equivalent plastic strain as ε̇pl = γ( σ σ0 − 1)1/m, (4) where γ is the material viscosity parameter, σ is the effective stress, m is the strain rate hardening coefficient, σ0 is the static yield point of the elasto-plastic model. two different nanoindentation machines were used and the tip geometry plays an important role during the fitting load-displacement data. thus, the parameters were identified only to best-fit the experimental ni curve (speed = 0.67 mn/s). the parameters were identify as: m = 0.8, γ = 0.025, density = 2000 kg/m3, e = 33.8 gpa, ν = 0.2, σ0 = 300 mpa, tangent modulus = 2000 mpa. after parameters were identified, the same set of parameters was used for xpm loading (speed = 20.0 mn/s). the results of fitting are shown in fig. 5. the vonmises equivalent strain under the tip was analyzed at the end of loading period, see fig 4. it was found the during fast loading, smaller penetration depths and smaller strain are reached compared to ni but shorter loading time is leading to higher 0 50 100 150 200 250 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 displacement [nm] lo ad [m n ] avg. exp.; speed = 0.67 mn/s model ni; speed = 0.67 mn/s model xpm; speed = 20.0 mn/s experimental data spread figure 5. calculated responses of ni and xpm loading with rate dependent viscoplastic model. strain-rate for xpm method. the results are the following: vonmises strain = 0.1542 (xpm) and 0.2009 (ni), strain rate = 1.542 s−1 (xpm) and 0.067 s−1 (ni). from the above, it is obvious that xpm loading leads to 23x higher strain rate comprared to ni. to quantify the effect of strain rate the oliver and pharr theory [2] was applied on model responses. it was found that e increased by 10 % for xpm mode. both experiment and numerical simulation show the shift of e to higher values. the small difference between numerical modeling and experiment could be given by overlaps of individual phases within the main peak. 5. conclusions this study shows mechanical characterization of portland cement paste with two nanoindentation testing modes: classical quasi-static grid nanoindentation (ni) and accelerated property mapping (xpm). both techniques provide qualitatively comparable results of elastic modulus and hardness. standard grid nanoindentation is favorable for covering larger areas and measuring creep parameters. xpm provides very fast acquisition time of hundreds indents in minutes with very high spatial resolution. the method is very good 99 j. němeček, j. němeček, j. lukeš acta polytechnica in identifying individual phases in a single measured map. also, xpm artificially shifts results to higher values due to the strain rate effect. this shift was quantified as 14.8 % from the experiment and calculated as 10 % according to numerical simulation. in this context, the oliver and pharr theory which is implemented as standard in nanoindenter’s softwares should be used with caution when strain-rate sensitive materials are tested with xpm technique. acknowledgements financial support of the czech science foundation (project 17-05360s) and the grant agency of the czech technical university in prague (sgs20/107/ohk1/2t/11 ) are gratefully acknowledged. references [1] a. c. fischer-cripps. nanoindentation. springer, 2002. doi:10.1007/978-0-387-22462-6. [2] w. c. oliver, g. m. pharr. an improved technique for determining hardness and elastic modulus using load and displacement sensing indentation experiments. journal of materials research 7(6):1564–1583, 1992. doi:10.1557/jmr.1992.1564. [3] e. d. hintsala, u. hangen, d. d. stauffer. high-throughput nanoindentation for statistical and spatial property determination. jom 70(4):494–503, 2018. doi:10.1007/s11837-018-2752-0. [4] p. s. phani, w. oliver. ultra high strain rate nanoindentation testing. materials 10(6):663, 2017. doi:10.3390/ma10060663. [5] p. h. bischoff, s. perry. compressive behaviour of concrete at high strain rates. materials and structures 24(6):425–450, 1991. doi:10.1007/bf02472016. [6] j. němeček, v. králík, j. vondřejc. micromechanical analysis of heterogeneous structural materials. cement and concrete composites 36(0):85 – 92, 2013. doi:10.1016/j.cemconcomp.2012.06.015. [7] g. constantinides, f.-j. ulm. the nanogranular nature of c–s–h. journal of the mechanics and physics of solids 55(1):64–90, 2007. doi:10.1016/j.jmps.2006.06.003. [8] i. n. sneddon. the relation between load and penetration in the axisymmetric boussinesq problem for a punch of arbitrary profile. international journal of engineering science 3(1):47–57, 1965. doi:10.1016/0020-7225(65)90019-4. [9] m. sebastiani, r. moscatelli, f. ridi, et al. high-resolution high-speed nanoindentation mapping of cement pastes: unravelling the effect of microstructure on the mechanical properties of hydrated phases. materials & design 97:372 – 380, 2016. doi:10.1016/j.matdes.2016.02.087. [10] p. perzyna. fundamental problems in viscoplasticity. vol. 9 of advances in applied mechanics, pp. 243 – 377. elsevier, 1966. doi:10.1016/s0065-2156(08)70009-7. 100 https://doi.org/10.1007/978-0-387-22462-6 https://doi.org/10.1557/jmr.1992.1564 https://doi.org/10.1007/s11837-018-2752-0 https://doi.org/10.3390/ma10060663 https://doi.org/10.1007/bf02472016 https://doi.org/10.1016/j.cemconcomp.2012.06.015 https://doi.org/10.1016/j.jmps.2006.06.003 https://doi.org/10.1016/0020-7225(65)90019-4 https://doi.org/10.1016/j.matdes.2016.02.087 https://doi.org/10.1016/s0065-2156(08)70009-7 acta polytechnica 27(0):96–100, 2020 1 introduction 2 materials and methods 2.1 sample preparation 2.2 nanoindentation 2.3 standard evaluation 3 results and discussion 3.1 standard nanoindentation 3.2 accelerated property mapping 3.3 indentation creep parameter 3.4 mutual comparison 4 numerical modeling 5 conclusions acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0032 acta polytechnica 27(0):32–36, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app local mechanical properties of atmospheric sprayed molybdenum coatings deposited with cascaded plasma torch jakub antoša, ∗, petra šulcováa, kateřina lencováa, šárka houdkováa, josef duliškoviča, jan hajšmanb, karolína burdováb, antonín račickýb a research and testing institute, tylova 1581/46 , 301 00 plzeň, czech republic b university of west bohemia, univerzitní 8, 306 14 plzeň, czech republic ∗ corresponding author: antos@vzuplzen.cz abstract. increasing interest for industrial use of thermally sprayed coatings leads to development in most thermal spraying technologies, including atmospheric plasma spraying (aps). the latest cascaded torch sinplexpro from oerlikon metco incorporates the efficiency advantages of cascaded arc technology into a single-cathode spray gun. it leads to more stable plasma arc across a wide range of gas flows, mixtures and pressures and also highly increases powder throughput. thermal sprayed molybdenum coatings are widely used for improving wear resistance and sliding properties in many mechanical applications. results of pure molybdenum coatings sprayed with cascaded plasma torch are not yet fully investigated. in this paper, mechanical properties of atmospheric plasma sprayed molybdenum coatings on steel (s235) substrate are evaluated. optimization of spraying parameters for spherical mo powder sprayed by cascaded plasma torch sinplexpro is carried out and the influence on final microstructure, mechanical and tribological properties of molybdenum coatings is analysed. keywords: atmospheric plasma spraying, cascaded plasma torch, mechanical and tribological properties, molybdenum coatings. 1. introduction atmospheric plasma spraying (aps) technology is widely used for spraying of function layers. plasma arc technology offers high temperature and fast and stable flow of plasma jet. high temperature is required for melting and consequential efficient deposition of refractory metals, ceramics and cermets. capability to deposit such materials makes plasma spraying particularly suitable for the production of wear resistant coatings, corrosion protective layers and thermal barriers [1]. molybdenum layers are used in many industry applications for its wear and scuff resistance and to reduce friction between sliding components [2]. whereas the spraying of pure molybdenum and molybdenum alloys by conventional plasma torches has been closely investigated in the past (e.g. ref. [2–4]), molybdenum spraying with cascaded plasma torch is not yet fully examined. this paper concerns spraying of pure spherical industrial grade molybdenum powder amdry 313x by oerlikon metco by cascaded plasma arc torch sinplexpro. correlation between spraying parameters and mechanical and structural properties of molybdenum coating on mild steel substrates are investegated and the optimization of spraying parameters of molybdenum powder for plasma torch sinplexpro is carried out. industrial aps with conventional torches is often struggling with reproducibility and reliability [5]. this is caused due to the instability of arc and subsequent oscilation of the output power level of the conventional torch during spraying process. sinplex pro cascaded torch deals with this issue by stacking copper rings insulated one from each other into a cascade. those rings are called neutrodes. the whole cascade of neutrodes is positioned between cathode and the nozzle, which serves as an anode. electric arc developing in such configuration is more stable and reduces almost all problems of conventional torches [5–7]. 2. setup of experiment in this experiment, argon is used as the primary gas and hydrogen as the secondary gas. argon and hydrogen are mixed to form plasma gas. this mixture runs through the electric arc, dissociating and ionizing into high temperature plasma current. rapid increse of temperature up to 14 000 k [1] is accompanied by substantional volume expansion. after plasma current leaves the torch, spraying material in form of a fine powder is injected into the plasma jet. argon is commonly used as the carrier gas for the powder. plasma jet transfer both thermal and kinetic energy into the particels. melted and accelerated particles are deposited on the substrate material [1]. in the first stage of design of this experiment, it was required to choose variable parametrs among all process parameters of spraying. main process parameters of sinplex pro system are as follows: diameter 32 http://dx.doi.org/10.14311/app.2020.27.0032 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 atmospheric sprayed molybdenum coatings mechanical properties of torch nozzle, plasma current, primary gas flow and secondary gas flow. these four parameters take critical role in forming of plasma jet they determine both temperature and plasma current velocity, thus directly affecting melting of the particles and their velocity. secondary process parametrs are spraying distance, relative movement speed of the torch over the sprayed specimen, mass flow of the powder and carrier gas flow. according to ref. [2, 3] concerning the influence of process variables, three critical parameters were determined: velocity of particles (vp), temperature of particles (tp) and temperature of substrate (ts). those three parameters have the most significant effect on the forming of the layer as well as its final properties. the state of particles described by velocity and temperature express its kinetic an thermal energy state. among the setup parameters, the electric current has the most determinating effect on the output power level and temperature of the plasma jet and thus directly affecting the particle temperature and melting. plasma gas flow rate as another main setup parameter has crucial effect on the velocity of the particles, and secondary effect on the plasma jet temperature. substrate temperature as the last critical parameter is a result of combined effects of plasma jet conditions, spraying pattern, torch movement speed, spraying distance, eventual cooling cycles and the dimensions and thermal properties of the substrate istelf. it should be noted that the previous statements about influence of main setup parameters (current, gas flow, spraying distance etc.) on the three critical parameters (vp, tp, ts) are simplified in order to briefely describe the main effect of each setup parametr. in reality all setup parameters take effect in a most complex manner as one dynamic system, simultaneously influencing each others as well as all ciritcal parameters. for this reason it is essential, in order to simplify the experiment, to reduce the number of variable parametrs. in the design stage of experiment it was decided to use 9 mm diameter torch nozzle and constant spraying distance of 125 mm. main reason for choosing the 9 mm nozzle (the largest diametr of commercially availible nozzles) is economical aspect. according to ref. [8], larger nozzle diametr increases spraying power output, improving the spraying process efficiency and leading to faster and thus more economical spraying times. the choosing of constant spraying distance may be considered questionable, because the impact of different spraying distances on the final properties of the sprayed layers could be substantial. as mentioned in ref. [4], increasing the spraying distance generally leads to higher hardness, which possitively correlates with wear ressistance but may not have possitive influence on the fritcion and sliding properties. this hardness growth happens due to the increased level of oxidized particles that is because of longer exposure time of melted particles to the ambient atmosphere. the amount of oxidized particles is again not only funciton of the spraying distance alone, but comes in a complex system with other input parametrs especially powder feed rate, grain size, plasma jet velocity and temperature etc. overall, in order to clarify the basic relations in setup parameters properties, it was decided to maintain constant spraying distance during the experiment [3–5]. another constant parameters was powder feed rate (63 grams per minute) and carrier gas flow rate (6 nlpm). also secondary plasma gas (hydrogen) flow rate was set on constant level for the whole experiment (5 nlpm). after setting most setup parameters to constant levels, electric current and primary gas flow rate was the only two parameters left to vary. table 1 displays the chosen combinations of those two parameters for the experiment. current varried between 350 a and 600 a in 50 a intervals, while primary gas flow rate between three levels 40, 50 and 60 nlpm. not all sets of those parameter were included. to decide which sets should be sprayed, a cpsp (critical plasma spraying parametr [9–11]) parameter was used to determine the thermal state of plasma jet for particular set of conditions. cpsp is defined as the ratio between the torch input power and primary plasma gas flow rate [10]. from all the possible combination of parameters, ten final parameter sets were chosen in order to fill the interval of cpsp between 0.5 and 1.4. all specimens were degreased and grid blasted shortly before the spraying process. spraying took place in a plasma spraying chamber of research and testing institute pilsen. conditions for all spraying were the same, using same batch of the powder, same spraying program and also all equipment remained the same for all spraying sets. temperature monitoring was conducted during spraying process in order to prevent overheating of the substrate and eventual cooling cycles were included. ten passes of the torch were used on all sets. 3. testing methods layer thickness, surface roughness and macro hardness hr15n were measured on all sprayed specimens. layer thickness was measured as a difference between specimen thickness before and after coating deposition process and cooling to the ambient temperature using micrometer screw gauge. the measured thickness was double checked using optical metallography measurement on the scratch pattern. scratch patterns of samples from all sets were inspected and photographed using optical microscopy. all samples also underwent scanning electron microscopy and eds analysis. adhesive strength of coating to the substrate was evaluated according to iso en 14916 on specimens from 3 chosen set of parameters. a major effort was devoted to the detailed measurement of hardness and modulus of molybdenum coatings. rockwell hardness test hr15n was employed 33 j. antoš, p. šulcová, k. lencová et al. acta polytechnica figure 1. microstructure of samples from selected sets (100x). input current primary gas fow rate ar 40 nlpm ar 50 nlpm ar 60 nlpm 350 a cpsp 0.5n. 1 400 a cpsp 0.6n. 2 450 a cpsp 1.1n. 9 cpsp 0.9 n. 7 cpsp 0.7 n. 3 500 a cpsp 0.8n. 4 550 a cpsp 0.9n. 5 600 a cpsp 1.4n. 10 cpsp 1.2 n. 8 cpsp 1 n. 6 table 1. matrix of chosen variable parameters for experiment. cells displays a serial number and an estimated value of cpsp for particular set of spraying parameters on the polished surface of each specimen. tests concerning vickers hardness included hv0.3 for random measurement on scratch pattern and indentation map of hv0.1, which was used to measure hardness as a function of distance from the surface of the coating. nanoindentation with berkovich tip for hardness measurement and for evaluation of modulus was carried out in two layers 150 microns underneath the surface and 150 microns above substrate. rather large indentation force of 5 n was required in order to obtain indents large enough to cover a representative area of coating including all typical coating formations. smaller indents would have measured only properties of splats, oxide layers or unmleted particles alone. hardness tests hr15n and hv0.3 were performed on specimens from all sets of parameters, whereas hv0.1, nanoindentation, indentation fracture toughness test (ift) and ball-on-flat tests were performed on three chosen sets. the three sets that underwent more detailed test were chosen according to cpsp values and microstructure in order to represent three typically obtained structures. ball-on-flat tribological test according to astm g-133 standard was performed on three chosen sets (n.4, n.5 and n.10). dry contact conditions, 25 n normal force, length of track 10 mm, 5 hz cycling frequency and 1000 s test duration was used. stainless steel ball of 62 hrc and 6.3 mm diameter was used against the polished surface of tested samples. three tests were performed on each specimen. the tracks left after this test were measured with a profilometer device to obtain volume loss values and underwent detailed sem screening and eds testing to clarify the tribological processes. 4. results and discussion 4.1. microstructure based on comparison of microstructure images, three sets were chosen for further testing: n.4, n.5 and n.10 (see 3 testing methods). images from optical microscopy of those three sets are shown above (fig. 1). 4.2. hardness and modulus results of hr15n surface measurement showed only small differences between individual sets, mostly in magnitude of standard deviation. hv0.3, hv0.1 as well as nanoindentation were performed on specimens of three chosen sets, with 10 random indents each. fig. 2. shows progression of hv0.1 as a function of distance from the coating surface. n.5 displayed a distinctive compact structure of bottom half of the coating (no pores were visible and splat boundaries were more difficult to recognize 34 vol. 27/2020 atmospheric sprayed molybdenum coatings mechanical properties figure 2. results of hv0.1 indentation as a function of distance from surface. using the optical microscopy). this compact layer is accompanied with an increase in hardness. specimens representing n.4 and n.10 lacks this compact layer. n.10 represents the most porous and coarse microstructre of the three chosen sets, while n.4 is finer, with smaller and more evenly located pores. tab. 2 displays the results of nanoindentation, where values are highlighted by colored columns. it showed that n.5 containing visible bottom compact layer has the highest modulus and hardness in both measured layers (below surface and above substrate). especially significant increase in modulus implies that n.5 has the largest residual compressive stress, especialy in the compact layer [11]. this may happen due to the rather large input power level combined with high flow rate of plasma gas. table 2. resulst of nanoindentation correlated to cpsp parametr 4.3. tribological properties results of ball-on-flat test performed on the three representative types showed two stages of dry friction contact (see fig. 3). in the first stage there was an increasing progress of coefficient of friction (cof). this initial stage lasted about 350 s for n.4 and n.5 and around 280 s for n.10. after this period of time, sudden drop in cof appeared and the progress remained constant and stable for the rest of the test. sem images with eds analyses implied, that the main mechanism for this cof decrease was creation of oxidized films. with the shortest initial period, n.10 also displayed lowest cof value of all tested specimens for both stages 0.64 in the first stage following a drop to 0.62. n.4 and n.5 cof values were 0.68 (dropping to 0.66) and 0.7 (dropping to 0.67) figure 3. comparison of cof progress of ball-onflat tested specimens. note: for illustration only one track for every specimen is displayed respectively. those results correspond to wear coefficient k, which was similar for n.4 and n.5 1.67e-4 and 1.91e-4 mm3/nm respectively. n.10 showing the lowest cof values also displayed significantly lower k, valuing 1.1e-4 mm3/nm it is 34% less to n.4 and 42% less to n.5. this difference in k value may be partly contributed to the difference in the duration of the first cof progression period, where presumably most material removal occurred. 5. conclusions results of experiment showed, that with increasing power input also grows mechanical properties hardness and modulus. higher plasma gas throughput has the similar effect. the increase in modulus can be a result of the increase in residual stress, with higher modulus indicating higher compressive stress [11]. as implied in ref. 12 that higher deposition temperature leads to increase in splat cohesion and thus higher modulus. higher deposition temperature also shifts the stress towards compression. all those aspects are already known and have been described in detail in the past. therefore these knowledges can be successfully applied to design mechanical properties of plasma sprayed molybdenum with cascaded torch as well. however, as the main industrial application for molybdenum coatings is wear resistance and sliding properties, the principal property for designing the coating should be also its final wear resistance and sliding properties. therefor sets of parameters represented by specimen n.10, although not having the highest hardness nor modulus, can be proclaimed as the best tested coating for the required industrial applications. structure of this coating with the best cof and wear resistance was achieved with high power level of the plasma jet of 600 a, and rather low primary argon flow rate of 40 nlpm. as a result of those parameters, an increase in oxidation level can be expected in this coating compared to others, but eds analyse did not indicate any relevant difference. 35 j. antoš, p. šulcová, k. lencová et al. acta polytechnica acknowledgements the paper has originated in the framework of project no. tj02000135. references [1] m. i. boulos, p. fauchais, j. heberlein, et al. thermal plasma technology. vol. 2. international union of pure and applied chemistry, 1999. [2] an integrated study of thermal spray process-structure-property correlations: a case study for plasma sprayed molybdenum coatings. materials science and engineering a, structural materials: properties, microstructure and processing 403, 2005. doi:10.1016/j.msea.2005.04.056. [3] s. sampath, x. jiang, a. kulkarni, et al. development of process maps for plasma spray: case study for molybdenum. materials science and engineering: a 348(1):54 – 66, 2003. doi:10.1016/s0921-5093(02)00642-1. [4] t. stolarski, s. tobe. the effect of spraying distance on wear resistance of molybdenum coatings. wear 249(12):1096 – 1102, 2001. doi:10.1016/s0043-1648(01)00842-0. [5] r. chidambaram seshadri, s. sampath. characteristics of conventional and cascaded arc plasma spray-deposited ceramic under standard and high-throughput conditions. journal of thermal spray technology 28(4):690–705, 2019. doi:10.1007/s11666-019-00841-9. [6] oerlikon metco. plasma spray gun sinplexpro: user manual. 2018. [7] an introduction to thermal spray. bro-0005.6 . oerlikon metco, july 2016. issue 6. [8] s. t. aruna, n. balaji, j. shedthi, v. k. w. grips. effect of critical plasma spray parameters on the microstructure, microhardness and wear and corrosion resistance of plasma sprayed alumina coatings. surface & coatings technology 2012. doi:10.1016/j.surfcoat.2012.08.016. [9] d. zhao, f. luo, w. zhou, d. zhu. effect of critical plasma spray parameter on complex permittivity and microstructure by plasma spraying cr/al2o3 coatings. applied surface science 2013. doi:10.1016/j.apsusc.2012.10.061. [10] s. yugeswaran, v. selvarajan, d. seo, k. ogawa. effect of critical plasma spray parameter on properties of hollow cathode plasma sprayed alumina coatings. surface and coatings technology 203(1):129 – 136, 2008. doi:10.1016/j.surfcoat.2008.08.030. [11] j. matejicek, s. sampath, d. gilmore, r. neiser. in situ measurement of residual stresses and elastic moduli in thermal sprayed coatings: part 2: processing effects on properties of mo coatings. acta materialia 51(3):873 – 885, 2003. doi:10.1016/s1359-6454(02)00477-9. 36 https://doi.org/10.1016/j.msea.2005.04.056 https://doi.org/10.1016/s0921-5093(02)00642-1 https://doi.org/10.1016/s0043-1648(01)00842-0 https://doi.org/10.1007/s11666-019-00841-9 https://doi.org/10.1016/j.surfcoat.2012.08.016 https://doi.org/10.1016/j.apsusc.2012.10.061 https://doi.org/10.1016/j.surfcoat.2008.08.030 https://doi.org/10.1016/s1359-6454(02)00477-9 acta polytechnica 27(0):32–36, 2020 1 introduction 2 setup of experiment 3 testing methods 4 results and discussion 4.1 microstructure 4.2 hardness and modulus 4.3 tribological properties 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0062 acta polytechnica ctu proceedings 22:62–66, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app influence of corundum as coarse aggregate in high-performance concrete on projectile impact resistance michal mára∗, zdeňka říhová, markéta kočová czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: michal.mara@fsv.cvut.cz abstract. the aim of this paper is focused on the influence of corundum as coarse aggregate in high-performance concrete on projectile impact resistance in terms of depth of penetration and size of the crater. based on the experimental programme, samples of high-performance steel-fibre-reinforced concrete with different types and quantities of coarse aggregates were designed and produced, to be subjected to impact loading in the form of projectile impact. damaged samples were scanned using a 3d scanner and their surface was evaluated. the goal of the paper was to find the optimum content of coarse aggregate in the mixture in order to improve its resistance against projectile impact. keywords: high-performance concrete, hpc, coarse aggregate, corundum, projectile impact, depth of penetration, crater surface, 3d scanning. 1. introduction in recent years, great emphasis is placed on durability, strength and safety of concrete structures, especially civil and military structures, which could be the target of deliberate or accidental shock loads. these structures, including objects of strategic importance, such as government buildings, reinforcements, defence shelters or other protective structures, must be resistant against this type of loads. these elements and their material are designed to withstand the possible extreme loads that can be caused by projectiles, fragments, explosions, cars or aircraft impact, earthquakes or terrorist attacks. one of the materials that have a high potential to use for these structures is high-performance concrete with a dispersed steel reinforcement (hpsfrc). this research is focused on improving the resistance of hpsfrc slabs to projectile impact, with aiming to improve the properties while maintaining the thickness and weight of the slabs for easy handling. from previous studies [1–3] it is known that the increasing amounts of steel fibres over a certain limit worsens the workability of fresh concrete and lowers of compressive strength and modulus. for this reason, this work is focused on improving hpsfrc by adding coarse aggregate. concrete resistance against projectile impact depends on many aspects. improved resistance against impact loads with respect to penetration depth and crater diameter can be achieved by reducing water/cement ratio and increasing compressive strength of concrete. the presence of dispersed steel fibres then decreases the diameter of the crater. for these reasons, the use of hpsfrc is the ideal solution for building structures that could be potentially exposed to impact loads. it is known that the conventional fibre-reinforced concrete (frc) with normal strength matrix has a high energy absorption capacity when hit by a projectile. however, studies show that hpsfrc has a much greater energy absorption capacity in both quasi-static and dynamic loading [4]. the resistance of the concrete against the impact of the projectile is also affected by the size, strength and hardness of the added aggregate. in order to achieve the highest compressive strength, the presence of coarse aggregate in hpsfrc is usually reduced or even eliminated in order to achieve the required homogeneity and density of the mixture. the antipenetration capabilities of coarse aggregate have been therefore almost neglected. however, in the projectile penetration experiments performed by zhang et al. [5], langberg and markeset [6] or wu et al. [7], they found that the penetration depth of the projectile does not decrease as long as the compressive strength of the concrete reaches a certain boundary. the optimal compressive strength for hpsfrc, from which the protective structures will be constructed, has been designed to be 90-150 mpa in terms of comprehensive considerations of effective protection and cost of production [4]. 2. material and mixtures as a coarse material for hpsfrc, an artificial white corundum was chosen (figure 1), mainly because of its very high hardness (mohs hardness scale number 9) and its impact resistance. an experimental study with uhpsfrc with added corundum as coarse aggregate (uhp-casfrc) was carried out by wu [4], who focused on uhp-casfrc samples with different contents and different corundum fractions. the sam62 http://dx.doi.org/10.14311/app.2019.22.0062 http://ojs.cvut.cz/ojs/index.php/app vol. 22/2019 influence of corundum as coarse aggregate. . . figure 1. corundum, fraction 3-5 mm. ples were tested by a projectile with impact velocity in the range of 510-850 m/s, an average mass of 341.2 g and a diameter of 25.3 mm. from the results, it was concluded that when the aggregate grain size increases, the concrete achieves higher compressive strengths. when the fraction was changed from 5-20 mm to 3545 mm, the compressive strength increased by 18%. in terms of depth of penetration (dop) of the projectile, it can be stated that when the aggregate grain size is increased from 0.5d (where d is the diameter of the projectile) to 1.5d and 2.5d its value decreases gradually. with increasing grain size of filler, the probability of a direct hit of coarse aggregate particles increases. therefore, it is possible in this way greatly enhance impact resistance of uhpsfrc, if coarse aggregate with high hardness and strength is used. the surface of the crater after the impact of the projectile decreases as the grain size of the added coarse aggregate increases. slightly better results are also obtained when the content of corundum in the mixture is increased. the average surface area of the crater with 45% corundum content was 12.5% smaller compared to the 30% corundum content [4]. within the experimental program, six types of samples were proposed one reference mixture and five other mixtures with a variable percentage of corundum (10%, 15%, 20%, 22.5% and 30%). all slabs were made in a thickness of 50 mm and a size of 400 × 300 mm. the steel fibres were 13 mm in length and 0.15 mm in diameter. the amount of added fibres for all samples was 1.5% by volume of the resulting composite. a dry prefabricated mixture of a multifunctional silicate composite [8] was used, to which the corundum was added. this prefabricated mixture comprises a cement matrix which forms a ground cement clinker with a specific surface area of more than 150 m2/kg. the dry prefabricated mixture contains the high-range water reducer, a defoaming agent, silica-fume, silica powder and fine sands with a maximal grain size of 1.2 mm. the number of individual components can be seen in table 1. the water to cement ratio of this mixture was 0.3. two kinds of samples were made with the prefabricated mixture with 10% and 15% content of corundum from the total volume of the composite. the corundum was used in a 3-5 mm fraction; its figure 2. capture the crater by using a 3d scanner. bulk density was 3850 kg/m3. the producer did not provide us with the granulometry of this fraction. for further corundum samples, a custom mixture was designed to meet optimized hpsfrc granulometry. as well as prefabricated mixture it contained silica powder, microsilica, a superplasticizer and a defoaming agent. the composition of this mixture can be seen in table 2. in this recipe, corundum replaced fine aggregates in 15%, 22.5% and 30% of the total aggregates content. the mixture uses a cement class 52.5, the water-to-cement ratio was designed to 0.3. 3. experimental programme besides the slabs, three beams of 160 × 40 × 40 mm from each mixture were made. for these samples was performed a three-point test for flexural strength. average strengths and bulk densities of materials are shown in table 3. the table is supplemented with the average values of hpsfrc samples made from a dry prefabricated mixture of multifunctional silicate composite without added coarse aggregate. prepared slabs of 400 × 300 × 50 mm were subjected to a projectile impact 7.62 × 39 in the shooting range. it is a 7.62 mm full-metal jacket with a soft lead core. the weight of this projectile was 8.04 g, the velocity of the projectile is around 710 m/s. in previous experiments, uhpsfrc slabs 50 mm thick without added coarse aggregate stopped this type of projectile, so it is assumed that the plates with coarse aggregate should also resist this type the projectile [9]. evaluation of the results was carried out by using the 3d scanner david, which employs two-picture photogrammetry (stereophotogrammetry). for each sample, the crater on the front side was captured using this 3d scanner (figure 2). the scanner was placed at a suitable distance and the correct resolution was set and the surface image of the board was scanned (figure 3). to capture the entire surface of the crater were taken pictures from four different angles. these four pictures were putted together by software david and created one final 3d picture of crater. the resulting surface was exported in .obj format and subsequently processed in autocad civil 3d software. then, in this program, from the files that 63 m. mára, z. říhová, m. kočová acta polytechnica ctu proceedings mass proportions component 1 binder 0.1 active form sio2, so called microsilica 0.25 silica powder (fraction d50 around 6µm) 1.6 silica sand with smooth granulometry(0.1 1.2 mm medium size grain d50 – around 500µm) 0.01 high range water reducer (superplasticizer) 0.001 defoaming agents table 1. components of the prefabricated mixture – proportions by weight. composition hpsfrc corundum corundum corundum corundum corundum 10% 15% 20% 22.5% 30% cement 1 additives 0.4 water 0.3 hrwr 0.025 fine sands 1.76 1.59 1.50 1.42 1.36 1.23 corundum 0 0.17 0.26 0.34 0.40 0.53 table 2. composition of high-performance concrete with corundum. figure 3. picture the surface of the board. figure 4. picture of crater surface in autocad civil 3d using point cloud. contained the 3d coordinates of the scanned craters, it created a point cloud (figure 4). the software then created a 3d surface and was able to measure the area of the crater (3d surface), diameter of the created crater and the penetration depth of the projectile. 4. results and discussion the resulting values of diameter and depth of penetration of the surface craters were obtained using autocad civil 3d are shown in tab. 4. the following graph (figure 5) shows that adding 10% and 15% corundum aggregates leads to a decreasing penetration depth of 23.4 mm to 20.7 mm and at a further increase of corundum content the dop even decrease slightly to 19.7 mm. for samples of designed hpsfrc mixture, with increasing amount of corundum is achieved to reduce a depth of penetration from 23.4 mm to 19.7 mm. the chart in figure 5 shows the dependence of the depth of penetration on the amount of corundum added. in general, it can be stated, that with the increasing amount of aggregate, the depth of penetration decreases, but the optimal composition of the mixture must be still found. if the grain curve is not filled with sufficient fine grain aggregate, it will occur to microcracks of the composite due to insufficient compaction of the mixture. the next chart (figure 6) shows the dependence of the diameter of the crater on the amount of corundum added. with the higher content of coarse aggregate in hpsfrc, the diameter of the crater formed by the impact of a projectile is decreasing. in the prefabricated mixture, when the corundum is increased from 10% to 20%, the crater diameter is reduced from 81 mm to 68.2 mm. on the contrary, when com64 vol. 22/2019 influence of corundum as coarse aggregate. . . mixture corundum compressive strength tensile strength density [kg/m3] [mpa] [mpa] [kg/m3] hpsfrc 0 158.0 25.60 2373 corundum 10% 119.5 164.6 30.19 2293 corundum 15% 181.1 164.2 30.86 2400 corundum 20% 239.0 173.3 28.16 2433 corundum 22.5% 296.4 163.5 31.25 2391 corundum 30% 375.0 170.4 26.40 2468 table 3. mechanical properties of high-performance concrete with corundum. samples dop diameter 2d surface 3d surface [mm] [mm] [mm2] [mm2] corundum 10% 20.7 81.0 5158 43095 corundum 15% 23.4 77.2 4688 39619 corundum 20% 19.7 68.2 3710 47967 corundum 22.5% 20.7 75.6 4506 40107 corundum 30% 19.7 66.7 3506 28384 table 4. crater dimensions of uhpfrc after projectile impact. figure 5. dependence of depth penetration on corundum content. figure 6. dependence of crater diameter on corundum content. 65 m. mára, z. říhová, m. kočová acta polytechnica ctu proceedings paring the mixture with 22.5% and 20% corundum, the crater diameter drops from 75.6 mm to 66.7 mm. other variations in the crater diameter are due to the diversity between prefabricated and a custom mixture. both mixtures have different amount of fine sands and corundum aggregate. that’s why the corundum sample has a 22.5% larger crater diameter than corundum 20%, even if there is coarser aggregate. 5. conclusions in terms of resistance to impact of the projectile, it has been found the addition of coarse corundum aggregate to high-performance steel-fibre-reinforced concrete (hpsfrc) has a positive effect on its impact resistance. the depth of penetration, as well as the diameter of the crater, depends on the compressive strength of the composite and the amount of aggregate added. the higher the compressive strength of the material and the coarser aggregate added to the hpsfrc, the better the impedance properties the material proves. these properties also depend on the quality of the added aggregate, especially on its strength and hardness. coarse aggregates have shown great potential to improve the resistance of hpsfrc, as the addition of a higher amount of corundum has improved the hpfsrc impact behaviour. the aim of this article was not to compare the benefits of corundum depending on the price. however, as literature shows, in larger structures the higher corundum price is defended in case of a significant improvement in the resistance to the projectile impact. acknowledgements this paper was prepared as a part of research supported by the ctu in prague [project no. sgs18/057/ohk1/1t/11]. the authors gratefully acknowledge the financial support from the ministry of industry of the czech republic [project no. fv10547]. references [1] p. máca, r. sovják, p. konvalinka. mix design of uhpfrc and its response to projectile impact. international journal of impact engineering 63:158–163, 2014. doi:10.1016/j.ijimpeng.2013.08.003. [2] v. corinaldesi, g. moriconi. mechanical and thermal evaluation of ultra high performance fiber reinforced concretes for engineering applications. construction and building materials 26:289–294, 2012. doi:10.1016/j.conbuildmat.2011.06.023. [3] r. yu, p. spiesz, h. brouwers. mix design and properties assessment of ultra-high performance fibre reinforced concrete (uhpfrc). cement and concrete research 56:29–39, 2014. doi:10.1016/j.cemconres.2013.11.002. [4] h. wu, q. fang, j. gong, et al. projectile impact resistance of corundum aggregated uhp-sfrc. international journal of impact engineering 84:38–53, 2015. doi:10.1016/j.ijimpeng.2015.05.007. [5] m.-h. zhang, v. shim, g. lu, c. chew. resistance of high-strength concrete to projectile impact. international journal of impact engineering 31:825–841, 2005. doi:10.1016/j.ijimpeng.2004.04.009. [6] h. langberg, g. markeset. high performance concrete-penetration resistance and material development. in proceedings of the ninth international symposium on interaction of the effects of munitions with structures, pp. 933–941. 1999. [7] h. wu, q. fang, x. chen, et al. projectile penetration of ultra-high performance cement based composites at 510–1320 m/s. construction and building materials 74:188–200, 2015. doi:10.1016/j.conbuildmat.2014.10.041. [8] k. kolář, z. bažantová, p. konvalinka. suchá prefabrikovaná směs silikátového kompozitu. patent, úřad průmyslového vlastnictví, 2015. [9] t. vavřiník. odolnost vysokohodnotného betonu proti nárazu projektilu. bachelor thesis, fakulta stavební, čvut v praze, 2012. 66 http://dx.doi.org/10.1016/j.ijimpeng.2013.08.003 http://dx.doi.org/10.1016/j.conbuildmat.2011.06.023 http://dx.doi.org/10.1016/j.cemconres.2013.11.002 http://dx.doi.org/10.1016/j.ijimpeng.2015.05.007 http://dx.doi.org/10.1016/j.ijimpeng.2004.04.009 http://dx.doi.org/10.1016/j.conbuildmat.2014.10.041 acta polytechnica ctu proceedings 22:62–66, 2019 1 introduction 2 material and mixtures 3 experimental programme 4 results and discussion 5 conclusions acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0136 acta polytechnica 27(0):136–140, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of two techniques for evaluating spherical indentation data aleš materna∗, petr haušild, jan ondráček czech technical university in prague, faculty of nuclear sciences and physical engineering, department of materials, trojanova 13, praha 2, czech republic ∗ corresponding author: ales.materna@fjfi.cvut.cz abstract. flow curves of 15kh2mfa, sv 08kh19n10g2b and 08kh18n10t steels used for fabrication of wwer-440 nuclear reactor pressure vessel and core internals were obtained using the automated ball indentation (abi) test technique and compared with flow curves evaluated from the same measured load-displacement data and widely used oliver-pharr method. differences in results obtained by both studied methods do not exceed 12 % and are attributed to the amount of material pile-up. keywords: automatic ball indentation, oliver-pharr method, pile-up, true stress-strain curve. 1. introduction perhaps the most important test of a material’s mechanical response is the tensile test. for small volume of material, from which the standard tensile test specimens cannot be manufactured, one can use an instrumented indentation. for estimation of a stress-strain curve from indentation data, a spherical indenter has advantage among others with e.g. conical or pyramidal shapes. contact angle of spherical indenter changes with loading (contrary to conical or pyramidal indenters), which allows determination of more than one point of the stress-strain curve. on the other hand, the shape of the ball limits the achieved strains to 0.2. at this limit value, both indent and ball diameters equal. there are numerous models available to reduce the spherical indentation data to uniaxial stress-strain behavior of the test material (e.g. [1–3]). they differ in the level of physical simplification and in the computational complexity. the accuracy of such models relies substantially on proper determination of the contact area between the indenter and test material. the simple model based on the well-known hertz law is suitable for low indentation forces, when the deformations under the indenter are elastic. a more advanced method was introduced by oliver and pharr (op) [4], which calculates the projected contact area from elastic contact stiffness at peak load. this method, however, can lead to underestimation of the true contact area for materials exhibiting pile-up [5, 6]. to account for the material pile-up, the partial unloading technique and set of iterative equations was introduced by haggag [7] in automatic ball indentation (abi) testing. after determination of the contact radius, two tabor’s expressions [8] can be used to obtain representative stress and strain. these indentation values should be equivalent to true stress-strain uniaxial ones. in this work, the op and the abi techniques are compared on data sets obtained during the standard abi test procedure with progressive indentation and intermediate partial unloadings. all materials involved in this study are used for manufacturing of pressure vessel and core internals of nuclear reactor of wwer-440 type. 2. tested materials three various steels present at wwer 440 nuclear power plant reactor were involved in this study: (i) chromium-molybdenum-vanadium low alloy 15kh2mfa steel, which is used as pressure vessel base metal, (ii) 19 chromium/10 nickel niobium-stabilized austenitic stainless steel sv 08kh19n10g2b, which is used as pressure vessel outer layer cladding, and, (iii) chromium-nickel titanium-stabilized austenitic stainless steel 08kh18n10t, which is used as main structural material for the reactor internals (namely the core barrel, the core basket and the block of guide tubes). 08kh18n10t steel was studied in two deformation states: (a) as manufactured, and, (b) colddeformed with the thickness reduction 20 %. 3. methods and analysis mechanical properties of all test materials were evaluated from load p displacement h curve recorded during a progressive ball indentation with multiple intermediate partial-unloadings at one location on specimen’s surface. as a basis for evaluation, tabor’s relations between the representative strain εr and the representative stress σr on left sides and the contact diameter d and the applied force p during spherical indentation on the right side were used: εr = 0.2 d d , (1) σr = 4p δπd2 , (2) 136 http://dx.doi.org/10.14311/app.2020.27.0136 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 two techniques of spherical indentation data evaluation where d is the indenter diameter, and δ is the constraint factor. it follows from equation (1) that strain up to a maximum value of 0.2 can be achieved. in that limit case the ball penetrates material surface to its full diameter (d = d). representative values εr and σr are regarded as material true plastic strain εp and true stress σt. during the instrumented indentation, contact diameter d is not directly measured and must be approximated from p h curve. there are numerous approaches to that and two of them were chosen in this work for comparison. 3.1. oliver-pharr method the most widely used method for determining elastic modulus from instrumented indentation data is oliver-pharr (op) method [4], which evaluates contact diameter d as intermediate result from the contact depth hc according to equations d = 2 √ hc (d − hc), (3) hc = hmax − ε pmax s . (4) here, pmax is the force at maximum indentation depth hmax, s=dp/dh|h = hmax the initial unloading slope, and, constant ε = 0.75 for spherical indenter. the constraint factor δ in equation (2) was set to 3 in this study. 3.2. haggag method haggag in his abi test procedure [7] uses tabor’s equations (1) and (2), iteratively computed plastic indentation diameter dp (after unloading) instead of d, and, the constraint factor δ in in equation (2) dependent on the stage of deformation beneath the indenter. the value of dp comes from the hertz’s classical theory and has the form dp = 3 √√√√√√0.5cd h2p + ( dp 2 )2 h2p + ( dp 2 )2 − hpd , (5) where the indentation depth after unloading hp is extrapolated from measured unloading data and c = 5.47 · p ( 1 e1 + 1 e2 ) , (6) where e1 and e2 are the young’s moduli for the indenter and test material, respectively. according to statistical analysis of many experimental data it was proposed by francis [8] that the constraint factor δ in equation 2 is a function of the indentation variable φ = εpe2 0.43σt . (7) contrary to francis’s work, the abi procedure uses slightly modified functions for the constraint in the form: δ =   1.12 φ ≤ 1 1.12 + τ ln (φ) 1 < φ ≤ 27 δmax φ > 27 (8) where δmax = 2.87αm, τ = (δmax 1.12)/ln(27) and αm is the constraint factor index whose value varies between 0.9 and 1.25 for various structural steels. since no experimental data were available for calibration, value αm = 1.1 in the middle of the interval was used. 4. experimental details 4.1. mechanical testing all abi tests were carried out at room temperature using inspekt 20 kn testing machine and custommade testing device equipped with the replaceable tungsten-carbide ball and with two-arm extensometer for the measuring of the ball displacements. two different ball diameters were used: 2.5 mm for testing of 15kh2mfa and sv 08kh19n10g2b steels, 1.5 mm for 08kh18n10t steel. indentation loading sequence for each test material is evident from figures 1-3. for comparison purposes, standard tensile tests at room temperature using inspekt 100 kn testing machine were performed for 15kh2mfa and sv 08kh19n10g2b steels. tensile test specimens of 6 mm diameter were oriented in circumferential direction of the pressure vessel that corresponds to the direction of the indentation. 4.2. confocal microscopy in order to determine the amount of pile-up, it was necessary to obtain the topography of the spherical indentation. the depth profile was measured with olympus ols5000-saf confocal microscope employing a 405 nm wavelength laser diode. figure 1. abi test load displacement curve for 15kh2mfa steel. 137 a. materna, p. haušild, j. ondráček acta polytechnica figure 2. abi test load displacement curve for sv 08kh19n10g2b steel. figure 3. abi test load displacement curve for undeformed a cold deformed 08kh18n10t steel. figure 4. stress/plastic strain curves for 15kh2mfa steel obtained by various techniques. 5. results and discussion measured load displacement curves are for 15kh2mfa steel plotted in figure 1, for sv 08kh19n10g2b steel in figure 2 and for two deformation states of 08kh18n10t steel in figure 3. evaluated points which form the true stress/plastic strain curves are for 15kh2mfa steel plotted in figure 4, for sv 08kh19n10g2b steel in figure 5 and for two deformation states of 08kh18n10t steel in figure 6. figure 5. true stress/plastic strain curves for sv 08kh19n10g2b steel obtained by various techniques. figure 6. true stress/plastic strain curves for undeformed a cold deformed 08kh18n10t steel. steels with higher hardness exhibit lower penetration depths at the same load level and the higher flow stresses at the same strain levels. it can be observed in figure 3 and figure 6 for 08kh18n10t steel hardened by the cold working. however, more important differences are between the curves in figures 4 6 evaluated using simple oliver-pharr method and more sophisticated iterative abi test method. except the cladding material, op leads to higher stresses and lower strains compared to the abi technique. as can be seen in figures 4 and 5, true stress-strain data points evaluated by abi iterative procedure are in good agreement with standard tensile curves and this agreement can even be improved by the correction of tensile data to specimens necking. one possible explanation for differences between the op and abi results lies in chosen standard values of constraint factors, which can be inaccurate for tested steels. alternatively, it could be attributed to the amount of material pile-up or sink-in around the indent. in case of material pile-up, op method underestimates the contact area which could lead to increase of estimated stresses. it was proved experimentally in figures 7 9, 138 vol. 27/2020 two techniques of spherical indentation data evaluation where cross-sectional profiles for all indents are plotted. whereas materials pile-up around the indents in 15kh2mfa (figure 7) and 08kh18n10t (figure 9) steels, surface around the indent in sv 08kh19n10g2b remains flat (figure 8). susceptibility to pile-up can also be observed in p h plots. for linear indenters (e.g. conical or pyramidal), high final-to-maximum depth ratio hf /hmax > 0.8 indicates material pile-up [5]. for non-linear spherical indenter, the hf /hmax criterion is not as straightforward, because the deformation state under the indenter depends on the depth of penetration. it starts in pure elastic regime with sink-in at small depths and ends in fully plastic regime with possible pile-up. from this point of view, 15kh2mfa and sv 08kh19n10g2b steels are tested at the similar h/d ratios with expected higher hf /hmax ratio for piling-up 15kh2mfa steel (compare figures 1 and 2). on the other hand, pile-up is typical for materials with higher elastic modulus-to-yield stress ratio e/σy, which is higher for sv 08kh19n10g2b comparing to 15kh2mfa. this discrepancy in observed effect of the yield stress on pile-up occurrence can be attributed to ability of sv 08kh19n10g2b cladding to workharden to much higher plastic strains comparing to 15kh2mfa steel (compare tensile curves in figures 4 and 5). figure 7. cross-sectional profile of the spherical indent in 15kh2mfa steel. 6. conclusions the instrumented spherical indentation was used for characterization of tensile properties of base metal (15kh2mfa steel) and cladding (sv 08kh19n10g2b steel) of wwer-440 reactor pressure vessel and of main structural material of reactor core internals in two deformation states (undeformed and colddeformed 08kh18n10t steel). the aim of the study was to evaluate true stress/strain curves of all studied materials and to compare two methods of their evaluation: more complex abi test iterative procedure proposed by haggag and the relatively simple and available method of oliver-pharr mainly used figure 8. cross-sectional profile of the spherical indent in sv 08kh19n10g2b steel. figure 9. cross-sectional profile of the spherical indent in undeformed a cold deformed 08kh18n10t steel. for determination of elastic modulus and hardness of materials. whereas tensile curves computed according to abi equations match standard uniaxial tensile curves for 15kh2mfa and sv 08kh19n10g2b steels, the op method predicts about 12 % higher stresses for 15kh2mfa steel and about 7 % and 11 % higher stresses for undeformed and cold-deformed 08kh18n10t steel, respectively. these higher stresses are probably connected to the material pile-up, which has increased the real contact area contrary to the hypothetical one evaluated using simplified elastic assumptions of the op model. in case of sv 08kh19n10g2b steel, no pile-up was observed and both methods led to similar results. acknowledgements this work was supported by the technology agency of the czech republic under the project no. th02020565 and by the european structural and investment funds under the project cz.02.1.01/0.0/0.0/16_019/0000778 (centre of advanced applied sciences). 139 a. materna, p. haušild, j. ondráček acta polytechnica references [1] h. francis. phenomenological analysis of plastic spherical indentation. journal of engineering materials and technology, transactions of the asme 98:272–281, 1976. doi:10.1115/1.3443378. [2] p. au, g. lucas, j. sheckherd, g. odette. flow property measurements from instrumented hardness tests. non-destructive evaluation in the nuclear industry pp. 597–610, 1980. [3] b. taljat, t. zacharia, f. kosel. new analytical procedure to determine stress-strain curve from spherical indentation data. international journal of solids and structures 35(33):4411 – 4426, 1998. doi:10.1016/s0020-7683(97)00249-7. [4] w. oliver, g. pharr. an improved technique for determining hardness and elastic modulus using load and displacement sensing indentation experiments. journal of materials research 7(6):1564–1583, 1992. doi:10.1557/jmr.1992.1564. [5] a. bolshakov, g. m. pharr. influences of pileup on the measurement of mechanical properties by load and depth sensing indentation techniques. journal of materials research 13(4):1049–1058, 1998. doi:10.1557/jmr.1998.0146. [6] b. taljat, g. pharr. development of pile-up during spherical indentation of elastic-plastic solids. international journal of solids and structures 41(14):3891 – 3904, 2004. doi:10.1016/j.ijsolstr.2004.02.033. [7] f. haggag. in situ measurement of mechanical properties using novel automated ball indentation system. astm special technical publication pp. 27–44, 1993. doi:10.1520/stp12719s. [8] d. tabor. the hardness of metals. clarendon press, uk, 1951. 140 https://doi.org/10.1115/1.3443378 https://doi.org/10.1016/s0020-7683(97)00249-7 https://doi.org/10.1557/jmr.1992.1564 https://doi.org/10.1557/jmr.1998.0146 https://doi.org/10.1016/j.ijsolstr.2004.02.033 https://doi.org/10.1520/stp12719s acta polytechnica 27(0):136–140, 2020 1 introduction 2 tested materials 3 methods and analysis 3.1 oliver-pharr method 3.2 haggag method 4 experimental details 4.1 mechanical testing 4.2 confocal microscopy 5 results and discussion 6 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2021.31.0036 acta polytechnica ctu proceedings 31:36–40, 2021 © 2021 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague possible changes in transport conception of international and long-distance railway connections relating the construction of new high-speed railway lines in the czech republic pavel purkart∗, david vodák, tomáš javořík, martin jacura czech technical university in prague, faculty of transportation sciences, department of transportation systems, horská 3, 128 03 prague 2, czech republic ∗ corresponding author: purkapav@fd.cvut.cz abstract. the aim of this article is to deal with the issue of high-speed railway lines (hsl) in the czech republic in response to the fast connections conception ("rychlá spojení" in czech, abbreviation "rs"). currently the rs1, rs2 and rs4 projects are the most watched during the preparation of the hsl network construction. the authors would like to focus on this issue from the wider point of view and to assess possible impact on the current lines of long-distance trains, the location of graphical timetable nodes and the way of the transport service on the area of the state. keywords: high-speed railway lines, czech republic, graphical timetable, international transportation, long-distance transportation, travel time, transport junction. 1. introduction currently there is an ongoing development of czech hsl network in response to the fast connections conception. there are not any hsl in the czech republic now. this situation is not, with respect to the sustainable development and the situation in the neighbouring states of the former western europe, acceptable, but this topic is being revived and preparation of selected rs sections is accelerating. the construction of new high-speed railway lines (hsl) in the czech republic is currently an actual topic. the fast connections ("rychlá spojení" in czech, abbreviation "rs") system has defined following main directions of the fast railway network [1]: • rs1 praha (prague) – brno – ostrava – [poland], • rs2 brno – [slovakia/austria], • rs3 praha – plzeň (pilsen) – [germany] (western direction – münchen), • rs4 praha – [germany] (northern direction – dresden) with possible branch (most – chomutov, žatec, louny), • rs5 praha – hradec králové/liberec – [poland]. we can see the fast connections conception in figure 1. currently the rs1, rs2 and rs4 projects are the most watched during the preparation of the hsl network construction. the aim of this article is to focus on this issue from the wider point of view and to assess possible impact on the current lines of long-distance trains, the location of graphical timetable nodes and the way of the transport service on the area of the state. according to the správa železnic (the main czech railway infrastructure manager), first sections of the rs network should be put into operation by the year 2028, which makes this issue very actual. 2. goals of the article this article is mainly focused on the impacts of the rs network and changes in transport planning, especially: • evaluation of rs conception, • possible line routing in the international transportation, • discussing possible impacts in national long-distance transportation. 3. solution of the issue 3.1. evaluation of rs conception according to the research team, the basic design of the rs network is consistent. rs conception considers the most important national and international transport flows. decrease of the travel times to praha (prague) is shown in the figure 2. t is logical to solve rs1, rs2 and rs4 branches in the first place, but we should not omit rs3 branch, because it connects the area of plzeň and southern germany. this connection represents a strong transport relation, but currently it is not certain what parameters and routing should this route have between plzeň and germany. it is alarming, considering that more than thirty years have passed since the fall of the iron curtain. the defined rs5 direction is also suitable for reflection. rs5 route could be extended to the south bohemia (with connection to austria), because this area is currently not fully connected to the rs network. goals of rs network construction of railway transportation development are essential: 36 https://doi.org/10.14311/app.2021.31.0036 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 31/2021 possible changes in transport conception figure 1. rs network (continuous lines represent planned rs connections, dashed lines represent possible future connections, dotted lines represent alternatives), source: [2]. figure 2. estimated travel times decrease after the rs network construction, compared to the situation in 2018 (120/ 90 min = travel time in 2018/ prospective travel time after the hsl construction or conventional network modernization), source: [2]. 37 p. purkart, d. vodák, t. javořík, m. jacura acta polytechnica ctu proceedings • increasing the importance of the czech railway network for international transport, • cutting the travel times between praha and all regional capitals to 120 minutes or lower, to achieve this goal rs lines should be used both by high-speed and conventional trains, • lowering traffic density on the overloaded section of railway network, so suburban and freight trains could use it. another important issue is the organization of railway traffic on new rs lines. currently railway traffic is ordinarily organised according to the graphical timetable. this approach should be also applied on the new rs lines. principles of graphical timetable can be mathematically described by using two simple equations. the edge equation has the following formula (1): th = n · 1 2 tt (1) the formula (2) is the circumferential equation:∑ th = n ∗ tt (2) above-stated variables mean: • tt – tact time (service interval), • th – time weight of the edge between two tact nodes, • n – natural number (depends on individual situation) [3]. both these equations can be easily explained in order to understand their real meaning. the edge equation shows that vehicles from one line are passing each other after one-half of the tact time. in case of public transport line with tact time 60 minutes the vehicles of opposite directions pass each other every 30 minutes. vehicles pass on the route at the place where the symmetry axis is. on single-track railway line, we have to build a passing point. on the lines with two or more tracks, vehicles can pass each other on the route. we can assume that the perfect travel time between two important transport nodes equals one-half of the tact time. circumferential equation is related to the term circle (close sequence of edges and nodes) from the graph theory. it says that total count of system travel time on the all edges should ideally equal a multiple of natural number of the tact time. in such case, strong relations between the lines in the nodes are ensured. 3.2. possible line routing in long-distance international transportation construction of the rs1 and rs2 sections is crucial for national transport relationships, because it can significantly help to overloaded sections of first and third transit railway corridor. we can also assume, that lines on hsl will be integrated to the international connections – hsl are used both by national and international lines. there are already international lines on the czech railway network. those lines have high quality vehicles, but their travel speed does not exceed 100 km/h, which is not sustainable in the context of today’s "distance reduction". selected lines are even bypassing czech republic, because longer route outside of the czech republic offers better travel time. construction of the rs network should improve this situation and increase the importance of the czech railway network for international transportation. possible conception of international lines is shown on the network graphics in the figure 3. following lines of international trains can create the core of transport services in the rs network scenario: • wien – brno – praha – dresden – berlin. this line should ideally reach systemic travel time four hours along the entire route, which means one hour between each city. this should provide optimal travel time between individual nodes. current travel time is eight hours – construction of rs1 and rs2 can cut down the travel times by one half. it is alarming, that current fastest connection wien – berlin is not routed through the area of the czech republic, but only through germany and austria. travel time of connection routed through the area of the czech republic is slightly longer – this clearly proves unsatisfactory condition of our railway network for international transportation, • warszawa – ostrava – brno – praha – plzeň – münchen. this line also contains section of the rs3. current travel time of long distance train lines is eleven hours, but these lines are also not routed through the area of the czech republic. at the same time, there is not an intention to build the rs3 in the parameters of hsl. reaching the travel time under eight hours can be considered as a success. we should reach following travel times for the optimal locations of graphical timetable nodes: warszawa – ostrava three hours, ostrava – praha two hours and praha – münchen 3 hours. the above-stated network is basic and it consists of two lines with international meaning. for its addition and full use of the rs network, we can also assume following lines: • košice – ostrava – brno – praha – plzeň – nürnberg, similar to line warszawa – münchen in east-west direction. it connects east part of slovakia, czech republic and germany. travel time through czech republic should not exceed 4 hours. travel time on the area of other states should not be higher in order to keep the international line attractive for passengers, 38 vol. 31/2021 possible changes in transport conception figure 3. network graphics of possible line routing in the international transportation. 39 p. purkart, d. vodák, t. javořík, m. jacura acta polytechnica ctu proceedings • budapest – bratislava – brno – praha – most/chomutov – karlovy vary is a line, which connects the capitals of the czech and slovak republic. it does not have to end in the praha, but it can go to the area of most (on the designed rs line) and further to the karlovy vary (on the conventional line), • it is appropriate to mention the line linz – české budějovice – praha – warszawa in case of construction of the rs5. international lines can create following peak intervals on the area of the czech republic (assuming that they will be operated with the 60 minutes’ interval): • interval 15 minutes praha – brno, • interval 30 minutes praha – ostrava, • interval 30 minutes praha – plzeň. with these operational parameters and sufficient length of the trains, the international lines can serve the national relations without any additional national lines. for example, there is 15 minutes’ interval on the relation praha – brno, which is close to the parallel graphical timetable of subway. it is apparent that praha will become an important railway node at an international level. its significance could be very close to nodes such as wien and berlin. 3.3. possible impacts on the national long-distance transportation the above-mentioned international lines are going to bring significant travel times reduction also in the most important national lines. there will be also additional lines with lower importance. international lines are going to affect national transportation, because many passengers, who are using current national lines, are going to use new connections routed on rs/hsl. construction of rs/hsl is going to bring key changes in the current line routing. we can shortly mention following impacts on current lines: • line ex1 praha – ostrava is going to lose its importance; important cities (pardubice and olomouc) are going to remain outside of the rs, therefore it will be necessary to solve their connection. the rs network is probably not going to be important for pardubice, but we have to evaluate it for olomouc, • line ex3 praha – brno is going to lose its importance similar to line ex1 praha – ostrava, • ine ex5 praha – dresden – berlin also loses necessity for connection of those cities, in ústí nad labem case it depends whether it is on rs route or not (we can consider stopping of the line wien – berlin). we definitely have to change the conception of the line r15 praha – ústí nad labem – most – karlovy vary – cheb. if a new line with hsl parameters is built, including a branch to the most area, this line, in its current form, loses its importance for connection with praha. it is purposeful to divide this line into three lines, for example: . praha – most – chomutov – karlovy vary – cheb with systemic travel time praha – karlovy vary in phase one maximally two hours, then maximally one and half hour (for fast connection praha – karlovy vary), . praha – ústí nad labem – teplice – litvínov, . děčín – ústí nad labem – teplice – bílina – most – chomutov. • line ex6 praha – plzeň is primarily going to be routed on rs line, if there is not any segment of conventional line remained. we have to question the importance of the line r16 praha – plzeň – klatovy, which is partly regional, but, thanks to the strong transport connection between klatovy area and praha, it can be divided into two lines in plzeň – selected trains from klatovy area may be routed on rs line from plzeň. 4. conclusion we can expect smaller or bigger impacts from building the rs network on almost every line, however our main goal should be total passenger increase in public transportation, balanced use of railway network, reduction of travel times between the cities and reduction of environmental damages from the individual car transportation. it is necessary to point out, that high-speed railway transportation has to be attractive and affordable. otherwise our goals cannot be reached. acknowledgements all the facts presented in this article are based on the results of the research on the ctu faculty of transportation sciences, department of transport systems. this research was supported by the grant agency of the czech technical university in praha, grant no. sgs18/150/ohk2/2t/16 railway track parameters for transportation service optimization and no. sgs20/138/ohk2/2t/16 design and optimal use of railway infrastructure parameters. references [1] vysokorychlostní tratě v čr. [online], praha: správa železnic, 2020 [cit. 2020-02-22]. dostupné z: https://www.szdc.cz/vrt/. [2] m. švehlík, m. pinkava, p. provazník. postup přípravy vrt v čr. vědeckotechnický sborník čd. 2018, 25(46), 16. issn issn 1214-9047. [3] k. baudyš. časové prvky jř. systémové jízdní doby: technologie železnižní dopravy. praha, 2014. 40 https://www.szdc.cz/vrt/ acta polytechnica ctu proceedings 31:36–40, 2021 1 introduction 2 goals of the article 3 solution of the issue 3.1 evaluation of rs conception 3.2 possible line routing in long-distance international transportation 3.3 possible impacts on the national long-distance transportation 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0117 acta polytechnica ctu proceedings 26:117–125, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app on optimum design of frame structures marek tybureca, ∗, jan zemana, c, martin kružíkb, c, didier henriond, e a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b czech technical university in prague, faculty of civil engineering, departement of physics, thákurova 7, 166 29 prague 6, czech republic c czech academy of sciences, institute of information theory and automation, department of decision-making theory, pod vodárenskou věží 4, 182 08 prague 8, czech republic d university of toulouse, laas-cnrs, 7 avenue du colonel roche, 31400 toulouse, france e czech technical university in prague, faculty of electrical engineering, department of control engineering, karlovo náměstí 13, 121 35 prague 2, czech republic ∗ corresponding author: marek.tyburec@fsv.cvut.cz abstract. optimization of frame structures is formulated as a non-convex optimization problem, which is currently solved to local optimality. in this contribution, we investigate four optimization approaches: (i) general non-linear optimization, (ii) optimality criteria method, (iii) non-linear semidefinite programming, and (iv) polynomial optimization. we show that polynomial optimization solves the frame structure optimization to global optimality by building the (moment-sums-of-squares) hierarchy of convex linear semidefinite programming problems, and it also provides guaranteed lower and upper bounds on optimal design. finally, we solve three sample optimization problems and conclude that the local optimization approaches may indeed converge to local optima, without any solution quality measure, or even to infeasible points. these issues are readily overcome by using polynomial optimization, which exhibits a finite convergence, at the prize of higher computational demands. keywords: frame structures, global optimum, polynomial optimization, semidefinite programming, topology optimization. 1. introduction designing economical, efficient, and sustainable structures represents a major challenge of the contemporary society. while structural engineers literally explore designs that must satisfy the requirements of limit states analysis, there is usually an infinite number of such designs. regardless of the properties of this feasible design space, it is required to select only one, the best design. this design quality is measured by an objective function, which usually approximates the expenses of production. greatly simplified, one such (most common) criterion considers maximization of structural stiffness (while the amount of available material is limited), or, equivalently, minimization of structural volume (while requiring a certain structural stiffness) [1]. among the structural optimization problems the greatest progress has been achieved so-far in optimizing (the cross-sectional areas of) truss structures [2]. this development can be attributed to convexity of the feasible design space [3, sections 1.3.5, 3.4.3, and 4.8], as the (axial) structural stiffness is an affine function of the cross-sectional areas. on the other hand, when the bending stiffness comes into a consideration, convexity does not hold in general. quite surprisingly, these problems are also much less studied, and to our knowledge only local optimization algorithms have yet been developed [4, 5]. in this contribution, we investigate the problem of optimum design of frame structures. in particular, we develop four different methods in section 2: (i) general non-linear optimization solved by the interior-point method of fmincon, (ii) the first-order optimality criteria (oc) method [1, 5], (iii) reformulation of the problem into a non-linear semidefinite program (nsdp) solved by penlab [6], and (iv) a suitably modified polynomial optmization (po) problem (iii) solved globally using polynomial optimization methods [7] and the mosek [8] optimizer. we show that the latter po approach generates guaranteed lower and upper bounds on the solution, providing a means of assessing the design quality. finally, section 3 introduces a set of three sample optimization problems to compare the optimization approaches. 2. solution techniques to frame optimization 2.1. problem statement let us consider the problem of optimum design of frame structures composed a finite number of nodes, nn, and of admissible elements, ne, defining the socalled ground structure [2]. the frame elements are attributed with non-negative, and thus possibly zero, given-shaped cross-sectional areas a. for simplicity, 117 https://doi.org/10.14311/app.2020.26.0117 https://ojs.cvut.cz/ojs/index.php/app m. tyburec, j. zeman, m. kružík, d. henrion acta polytechnica ctu proceedings ki(ai) =   eiai `i 0 0 −eiai `i 0 0 12eiii `3 i 6eiii `2 i 0 −12eiii `3 i 6eiii `2 i 4eiii `i 0 −6eiii `2 i 2eiii `i sym. eiai `i 0 0 12eiii `3 i −6eiii `2 i 4eiii `i   (3) we assume, in the following text, that the moments of inertia of individual cross-sections ii are polynomials of degree two1, i.e., ii = ci,ia2i for some ci,i > 0, which includes all cross-sections with given aspect ratios of all their components. note that in such a case, each cross-section is fully determined by its area. optimizing the cross-sectional areas a we search the maximum stiff structures within the available volume v of a linear-elastic material. the structural stiffness is measured (inversely) by the compliance c, work done by external forces, f (a), c(a) = f (a)tu = utk(a)u, (1) where u constitutes the generalized displacement vector, and k(a) is the symmetric positive semi-definite stiffness matrix— a polynomial function of a, k(a) = ne∑ i=1 ki(ai), (2) assembled from contributions of individual elements ki(ai). the lower the compliance, the stiffer is the structure with respect to the external forces. in this contribution, we use the element stiffness matrix of euler–bernoulli frame elements (3), with ei denoting the young modulus. in (1), f (a) is the external force column vector—a linear function of a—to allow for self-weight, assembled from the contributions of elements f (ai), f (a) = ne∑ i=1 fi(ai). (4) in the following text, we assume that ∀a > 0 : k(a) � 0, i.e., the structure is not a kinematic mechanism and the stiffness matrix is positive definite (which is denoted by “� 0”) for all positive cross-sectional areas. since k(a) has therefore the full rank, we also have f (a) ∈ im (k(a)), where im(•) is the image of •. 1the same procedure can be employed for higher-degree polynomials, so that all cross-sectional parameters can be optimized concurrently. we restrict ourselves, however, to the polynomials of degree two to maintain a simpler notation. this optimization problem is formalized as min a,u f (a)tu (5a) s.t. k(a)u = f (a), (5b) `ta ≤ v , (5c) a ≥ 0, (5d) with ` being the column vector of the frame elements lengths. notice that in general, (5) constitutes a nonconvex non-linear optimization problem because of the bilinear objective function (5a) and the polynomial equilibrium equality (5b) with possibly singular k(a). on the other hand, the volume constraint (5c) and the cross-sectional areas non-negativity constraint (5d) are affine functions of the design variables, u and a, and are in turn convex. the optimization problem (5) can be solved to local optimality using standard numerical optimization techniques. in particular, we will solve this problem using the interior-point method implemented in the fmincon function of matlab. 2.2. optimality criteria the inherent difficulty of singularity of k(a) in (5b) can be circumvented by assuming a small positive lower-bound on the cross-sectional areas [1], which is denoted by ε in this study. consequently, the former topology optimization problem (5) is effectively transformed into the sizing one and the displacement field u can be excluded from the design variables2. notice that in the limit when ε → 0, these problems are equivalent, but the smaller ε the higher the condition number of k(a), and thus it is more difficult to solve (5b). hence, we assume ε = 10−6 in this study. considering (5) with ε1 ≤ a, its lagrangian function reads as l(a, u,λ,µ,ν) = f (a)tu + λt [f (a) − k(a)u] +µ ( `ta −v ) + νt (ε1 − a) , (6) with the lagrange multipliers λ, µ, and ν; and 1 denoting the vector of all ones. in addition to the primal feasibility (5b)–(5d), the karush–kuhn–tucker conditions require feasibility of the dual and complementary 2at the price of solving the equilibrium equation in each iteration. 118 vol. 26/2020 on optimum design of frame structures slackness, µ ≥ 0, (7a) µ ( `ta −v ) = 0, (7b) ν ≥ 0, (7c) νt (ε1 − a) = 0, (7d) together with the stationarity of the lagrangian with respect to u, 0 = dl(a, u,λ,µ,ν) du = [f (a) − k(a)λ]t , (8) and a. however, because the equation k(a)λ = f (a) in (8) possesses a unique solution due to k(a) � 0, we have λ = u. using this observation, the necessary first-order optimality conditions read as 0 = ∂l ∂ai = 2ut ∂f (a) ∂ai − ut ∂k(a) ∂ai u + µ`i −νi. (9) conditions (9) and (7d) then imply that at the optimum, the frame elements with ai > ε have equal constant energy µ = 1 `i ut ∂k(a) ∂ai u − 2 `i ut ∂f (a) ∂ai . (10) aiming to satisfy (10), optimality criteria methods [1] build update schemes which increase stiffnesses of elements with energies higher than µ, and, conversely, decrease stiffnesses of elements with the energy lower than µ. the levels are balanced in an iterative process, based on the value of µ. in each iteration, the relative change of the design variables is bounded by the move limit ζ, assumed as ζ = 0.2 in this paper. consequently, the fix-point update scheme reads as a (k+1) i = max { max { (1 − ζ)a(k)i ,ε } , a (k) i [ b (k) i ]η} (11) with the tuning parameter η = 0.3 and with b (k) i = ut ∂k(a) ∂ai u − 2ut ∂f (a) ∂ai µ`i . (12) clearly, if ∀i ∈ {1, . . . ,ne} : b (k) i = 1, we reach a local minimum as (10) is satisfied. combination of (11), (12) with (5c) allows us to write the (current) volume v = `ta(k+1) as a continuous function of the multiplier µ. it can be seen from (12) that v is a non-increasing function of µ. in fact, strict decreasing occurs when εi < a. consequently, the bisection algorithm is used to find µ such that the volume constraint is satisfied. 2.3. nonlinear semidefinite programming in this section, we describe another approach to eliminate the displacement field variables u from the optimization problem formulation. first, let us rewrite (5) as min a,c c (13a) s.t. c− f (a)tk(a)†f (a) = 0, (13b) `ta ≤ v , (13c) a ≥ 0, (13d) where k(a)† denotes the moore-penrose pseudoinverse of k(a). because we require that ∀a > 0 : k(a) � 0, the (possible) singularity of k(a) is caused exclusively by zero rows and columns belonging to the degrees of freedom without any attached finite element (or, equivalently, ai = 0 for all attached elements in that node). this assumption allows us to partition the stiffness matrix into a positive definite principal submatrix k̂(a) and zero blocks, so that the pseudo-inverse equals k(a)† = ( k̂(a)−1 0 0t 0 ) . (14) using the same partitioning, the force vector f (a) is split into ( f̂ (a)t 0t )t , in which the term 0t appears due to the original assumption that f (a) ∈ im(k(a)). consequently, we see that (13b) can be rewritten to c− f̂ (a)tk̂(a)−1f̂ (a) = 0, (15) and (13) is therefore equivalent to (5). from (14) we have k†(a) � 0, so that c ≥ 0 based on (13b). moreover, iff f̂ (a) 6= 0, we have both f (a)tk(a)†f (a) > 0 and c > 0. because we minimize c (13a) and f (a)tk(a)†f (a) is bounded from below, (13b) can be simplified to the one-sided inequality c− f (a)tk(a)†f (a) ≥ 0. (16) to use the generalized schur complement lemma, e.g., [9, theorem 16.1], we further need to show that[ i − k(a)k(a)† ] f (a) = 0 (17) holds, with i denoting the identity matrix. indeed, (17) is always satisfied because f (a) ∈ im(k(a)), so that we can substitute f (a) by k(a)v, where v is a vector of coefficients of the linear combination. then, (17) is equivalent to[ k(a) − k(a)k(a)†k(a) ] v = 0, (18) with the term in the square brackets always zero [9, lemma 14.1], as k(a) is symmetric. finally, application of the generalized schur complement lemma to (16) and (18) provides us with min a,c c (19a) s.t. ( c −f (a)t −f (a)t k(a) ) � 0, (19b) `ta ≤ v , (19c) a ≥ 0, (19d) 119 m. tyburec, j. zeman, m. kružík, d. henrion acta polytechnica ctu proceedings which is a non-linear semidefinite program equivalent to both (13) and (5). moreover, if we substitute c with 1/s, where 0 < s < ∞ is a measure of stiffness, (19) is, after the application of the (standard) schur complement lemma, e.g., [9, proposition 16.1], reducible to max a,s s (20a) s.t. k(a) −sf (a)f (a)t � 0, (20b) `ta ≤ v , (20c) a ≥ 0, (20d) which is a useful reformulation when f is constant, i.e., self-weight is not considered. it shall be noted that due to the polynomial matrix inequalities (19b) and (20b) both the optimization problems are non-convex in general. they can still be solved efficiently (to local optimality) using, e.g., augmented lagrangian methods [6, 10]. in this contribution, we solve these problems using the open-source penlab optimizer [6]. 2.4. polynomial optimization having introduced three different local approaches to frame structure optimization, it is natural and expected that one asks for a global optimization technique. in this section, we exploit the fact that although (19) is non-convex it is indeed a polynomial optimization (po) problem at the same time, which allows us to employ modern po techniques, namely the lasserre hierarchy [7, 11], successively building tighter and tighter convex outer semidefinite programming (sdp) approximations called relaxations [7, corollary 4.3]. to develop a more efficient formulation suitable for po, we first recognize that the design variables in (19) are all bounded both from below and above: 0≤ai≤ v `i , ∀i ∈{1, . . . ,ne}, (21a) 0≤c ≤ ĉ. (21b) while the lower bound in (21a) is caused by the nonnegativity of the cross-sectional areas (19d), the upper bounds arise from the volume constraint (19c): none of the structural elements can occupy larger volume than v . in the compliance case (21b), the lower bound3 is due to k(a) � 0, recall the discussion in section 2.3, and the upper bound is provided by an arbitrary feasible design, e.g., the uniform cross-sectional areas a = v 1t` 1, (22) which are used for the upper-bound computation in this paper. then, the compliance bound is computed from ĉ = f (a)tk(a)†f (a), (23) 3in fact the strict inequality 0 < c is satisfied in all nontrivial optimization problems. where k(a)† reduces to k(a)−1 in the case of (22). to improve numerical stability of the solution process, we state the optimization problem in terms of the scaled cross-sectional areas asc and scaled compliance csc instead of the original a and c, where the scaled variables are defined in the [−1, 1] domain. therefore, we have ai = v (asc,i + 1) 2`i (24) and c = ĉ(csc + 1) 2 . (25) rewriting (21) in terms of asc,i and csc as unit ball constraints a2sc,i ≤ 1, ∀i ∈{1, . . . ,ne}, (26a) c2sc ≤ 1, (26b) we finally arrive at an equivalent formulation to (19), min asc,csc 0.5ĉ(csc + 1) (27a) s.t. ( 0.5ĉ(csc + 1) −f (asc)t −f (asc)t k(asc) ) � 0, (27b) 2 −ne − 1tasc ≥ 0, (27c) 1 −a2sc,i ≥ 0, (27d) 1 − c2sc ≥ 0. (27e) in (27), the objective function (27a) and also the constraints (27b)–(27e) are all polynomial inequalities4 non-negative in the feasible region of (27), which is therefore a semi-algebraic set. 4the entries in the pmi (27b) are polynomials, so that the feasible region of the pmi is a semi-algebraic set. alternatively, we may require the roots of the characteristic polynomial of (27b) to be non-negative. we refer to [12] for more details. 120 vol. 26/2020 on optimum design of frame structures the optimization problem (27) can readily be solved by building the (lasserre) hierarchy of convex linear sdp relaxations, with a monotonously converging objective functions to the global optimum. the hierarchy is generated by the gloptipoly [13] software package interfaced with the yalmip [14] toolbox, and the underlying sdp relaxations are solved by the state-of-the-art mosek optimizer [8]. 2.4.1. solution process to simplify the notation, let us now denote the objective function polynomial (27a) by p0, and the constraining polynomials (27c) to (27e) by p1 to p3, respectively. further, let p4 denote the pmi (27b) and let x = ( csc asc,1 . . . asc,ne )t (28) be the vector of the design variables. moreover, let br(x) = ( 1 x1 x2 . . . xne+1 x 2 1 x1x2 . . . x1xne+1 x 2 2 x2x3 . . . x 2 ne+1 . . . xr1 . . . x r ne+1 )t (29) denote the polynomial space basis of the maximum degree r. then, we can express each of the polynomials pj,j ∈{0, . . . , 3} as a linear combination pj(x) = |br (x)|∑ β=1 pα,j,β(xα)β (30) of monomials xα = ne+1∏ m=1 xαmm , ne+1∑ m=1 αm ≤ r, (31) indexed in the basis br(x). the vectors α associate a non-negative integer number with each element in x, and the vector pα,j contains coefficients of the linear combinations of the monomials. notice that∑ne+1 m=1 αm ≤ dj, where dj stands for the degree of the polynomial pj. a similar approach is also applied to the elements of p4 [12]. further, introducing y, with its components yβ corresponding to a monomial in the basis br(x), we build the lasserre hierarchy of convex linear semidefinite programming relaxations min y |br (x)|∑ β=1 pα,0,βyβ (32a) s.t. mr(y) � 0, (32b) mr−dj (y) � 0, ∀j ∈{1, . . . , 4}, (32c) and solve it successively with an increasing relaxation order r until the globally optimal solution(s) are found. for all our test cases, this convergence was always finite. in (32), the matrix mr(y) is the moment matrix of the r-th order, and mr−dj is the (r − dj)-order localization matrix associated with pj or pj. for more details about these matrices, we refer the reader to [7, 12]. 2.4.2. recognizing global optimality there are (at least) two ways to recognize whether y∗r, a r-th order relaxation solution to (32), is globally optimal. the first (common) way is based a sufficient condition for global optimality of (32) [15, 16], rank mr(y∗r) = rank mr−d(y ∗ r), (33) in which y∗ constitutes the optimal solution in the basis br(csc, asc), and rank mr(y∗) is less or equal to the number of these solutions [7]. the values of the original design variables can be extracted using cholesky or singular value decompositions [16]. for the optimization problem (27), we introduce another sufficient condition for global optimality. let ãr be the (optimized) values of the cross-sectional areas and let c̃r denote the compliance found in the r-th order relaxation. indeed, c̃r ≤ c∗ [7]. moreover, because any feasible design a generates an upper bound ĉr, recall eq. (23), we have c̃r ≤ c∗ ≤ ĉr. (34) a globally optimal solution to (27) is found if there is no gap, ĉr − c̃r −→ 0. (35) note that (35) is substantially simpler to check than (33). if (35) is not satisfied in the (current) relaxation order r, we have at least the quality measure for the current solution in hand. 3. sample problems in this section, we describe and solve three rather small-scaled structural optimization problems, and compare the four optimization approaches. 3.1. cantilever beam consider a (simple) cantilever beam design problem, as shown in fig. 1a. the beam is discretized into ne finite elements of equal lengths, each of them assigned a prismatic square cross-section. the beam is loaded with a skew force at its tip. further, we assume (dimensionless) e = 1 and v = 0.1. it can be seen from table 1 that the problem is relatively “easy” to solve, as all of the tested algorithms converge to the unique global optima (shown in 1b–1e), proved by the po approach (figs. 1h–1j). in fact, we include this optimization problem mainly to show that po scales unambiguously worse with the problem dimension than the other methods, and that the optimality criteria (oc) method is the fastest one. using oc, we can optimize the discretizations of 150 or 300 elements in 0.4 and 0.8 seconds (figs. 1f and 1h). 121 m. tyburec, j. zeman, m. kružík, d. henrion acta polytechnica ctu proceedings t [s] ĉ c̃ a1 fmincon 0.7 107.50 0.100 oc 0.1 107.50 0.100 nsdp 0.2 107.50 0.100 po(1) 0.3 107.50 107.50 0.100 (a) 1 element t [s] ĉ c̃ a1 a2 a3 fmincon 0.6 80.30 0.142 0.102 0.056 oc 0.1 80.30 0.142 0.102 0.056 nsdp 0.3 80.30 0.142 0.102 0.056 po(1) 0.3 80.72 35.81 0.147 0.097 0.056 po(2) 0.3 80.30 80.30 0.142 0.102 0.056 (b) 3 elements t [s] ĉ c̃ a1 a2 a3 a4 a5 fmincon 0.4 77.19 0.151 0.128 0.103 0.075 0.043 oc 0.1 77.19 0.151 0.128 0.103 0.075 0.043 nsdp 0.5 77.19 0.151 0.128 0.103 0.075 0.043 po(1) 0.3 79.09 24.72 0.151 0.123 0.096 0.073 0.058 po(2) 0.7 77.37 76.34 0.154 0.131 0.103 0.072 0.040 po(3) 26.8 77.19 77.19 0.151 0.128 0.103 0.075 0.043 (c) 5 elements t [s] ĉ c̃ a1 a2 a3 a4 a5 a6 a7 fmincon 0.5 76.23 0.155 0.139 0.122 0.104 0.084 0.061 0.036 oc 0.1 76.23 0.155 0.139 0.122 0.104 0.084 0.061 0.036 nsdp 0.4 76.23 0.155 0.139 0.122 0.104 0.084 0.061 0.036 po(1) 0.3 79.81 20.00 0.149 0.130 0.112 0.096 0.081 0.069 0.062 po(2) 2.5 77.02 71.69 0.166 0.145 0.122 0.099 0.077 0.055 0.036 po(3) 550.1 76.23 76.23 0.155 0.139 0.122 0.104 0.084 0.061 0.036 (d) 7 elements table 1. comparison of the four optimization approaches on the design of the cantilever beam problem. bold text denotes the proven global optimum. 1 ne 1 ne 1 − 2 ne 1 1 2 ne nn 1 ne 1 30◦x y (a) (b) c∗ = 107.50 (c) c∗ = 80.30 (d) c∗ = 77.19 (e) c∗ = 76.23 (f) c = 75.19 (g) c = 75.19 1 2 0 20 40 60 80 100 120 relaxation order r c om pl ia nc e c ĉr c̃r c∗ (h) 1 2 3 0 20 40 60 80 100 120 relaxation order r c om pl ia nc e c ĉr c̃r c∗ (i) 1 2 3 0 20 40 60 80 100 120 relaxation order r c om pl ia nc e c ĉr c̃r c∗ (j) figure 1. cantilever beam design problem. (a) shows the design domain, (b)–(e) are the optimal topologies for 1, 3, 5, and 7 elements; (f) and (g) display optimized topologies computed by oc with discretization by 150 and 300 elements. figures (h)–(j) show the convergance of po for 3, 5, and 7 elements. 122 vol. 26/2020 on optimum design of frame structures 1 2 x y 1 1 1 1 2 3 4 5 6 7 8 9 10 1 2 3 4 5 6 (a) (b) c = 1042.20 (c) c∗ = 959.32 1 2 0 500 1,000 1,500 2,000 relaxation order r c om pl ia nc e c ĉr c̃r c∗ (d) figure 2. 10-beam structure design problem. (a) shows the design domain. while all the tested local optimization algorithms converged to the topology in (b), the global optimum design (c) obtained by po possesses a clearly different topology. figure (d) shows the convergence of po. t [s] ĉ c̃ a1 a3 a4 a5 a6 a7 a8 a9 a10 fmincon 1.0 1042.14 0.144 0.040 0.018 0.000 0.003 0.000 0.210 0.039 0.038 oc 0.6 1042.33 0.146 0.039 0.017 0.000 0.002 0.000 0.212 0.039 0.037 nsdp 1.0 1042.20 0.144 0.040 0.018 0.000 0.003 0.000 0.210 0.039 0.038 po(1) 0.3 1429.31 443.20 0.103 0.082 0.000 0.029 0.000 0.000 0.121 0.095 0.069 po(2) 5.2 959.32 959.31 0.070 0.186 0.000 0.043 0.000 0.098 0.000 0.064 0.000 table 2. comparison of the four optimization approaches on the design of 10-beam structure. bold text denotes proven global optimum. the cross-sectional area a2 is zero in all test cases. 1 2 3 4 5 1 1 2 3 4 5 6 2 2 2 2 2 10 (a) tp 5tp 10 t p (b) 1 2 3 0 500 1,000 1,500 2,000 2,500 relaxation order r c om pl ia nc e c ĉr c̃r c∗ (c) figure 3. girder beam design problem. (a) shows the design domain, (b) the cross-sectional shape, and (c) the convergence of polynomial optimization. t [s] ĉ c̃ a1 a2 a3 a4 a5 fmincon 0.4 0.394 0.256 0.000 0.003 0.000 oc 0.1 1372.25 0.010 0.017 0.022 0.025 0.026 nsdp 2.0 0.000 0.000 0.000 0.000 0.000 po(1) 0.2 1456.75 297.34 0.007 0.015 0.022 0.027 0.029 po(2) 1.0 1426.05 1286.44 0.007 0.016 0.023 0.026 0.028 po(3) 37.7 1372.25 1372.25 0.010 0.017 0.022 0.025 0.026 table 3. comparison of the four optimization approaches on the design of the girder beam problem. bold text denotes the proven global optimum. 123 m. tyburec, j. zeman, m. kružík, d. henrion acta polytechnica ctu proceedings 3.2. 10-beam frame structure second, we consider the topology optimization problem of designing circular cross-sections of the 10-beam structure shown in fig. 2a. this structure is loaded by two moments of magnitudes 1 and 2, placed at the nodes 2 and 3 , respectively. note that there is no intersection between the elements 3 and 4 , and between 6 and 7 . we assume the young modulus e = 1 and the volume bound v = 0.5. evaluating the four optimization methods, table 2, we see that while the local optimization methods converged to nearby local minima of topologies shown in fig. 2b, po found the global optimum of a clearly different topology, fig. 2c. moreover, the global optimum possesses the compliance lower by 8%, which was reached in the second relaxation, fig. 2d. 3.3. i-shaped girder with self-weight last, we consider a simply-supported i-shaped girder beam of the span 20, loaded by a uniform load 1 and self-weight. due to the problem symmetry, we consider only one half of the problem, fig. 3a, and discretize it using 5 finite elements of equal length. this (half of the) girder beam is allowed to utilize at most v = 0.2 material of the young modulus e = 104 and of density ρ = 3. the beam cross-sections are parameterized by the parameter tp, which denotes the thicknesses of the flanges and web. the cross-sectional area equals a(tp) = 18t2p. we assume that the upper surfaces of the top flanges are aligned, so that we have i(tp) = 696t4p = 58/27a(tp)2. consequently, we optimize ai of individual elements, rather than tp. table 3 reveals that for this specific problem the oc converged to the global optimum. on the other hand, the fmincon and non-linear semidefinite programming approaches converged to an infeasible point. polynomial optimization exhibited convergence to the global optimum in three relaxations, fig. 3c. notice, moreover, that the designs found for r = {1, 2} were of very high qualities. 4. conclusions in this contribution, we investigated four topology optimization techniques for the design of frame structures with a given aspect-ratios all components of their cross-sections. three of the optimization techniques— general non-linear formulation solved by the fmincon solver, optimality criteria, and non-linear semidefinite programming—provide a local solution to the considered optimization problem, and in turn converge to a local optimum in general. unfortunately, the nsdp and fmincon techniques also converged to infeasible points. from the local optimization approaches, the optimality criteria method seems to be the most efficient one. on the other hand, the last technique — polynomial optimization — allows to solve the problem to proven global optimality, while generating both lower and upper bounds in each of the relaxation order. compared to the local approaches, the designs obtained in lower relaxation orders are of a known quality with respect to the optimum. however, finding a proven global optimum requires a considerably higher computational resources than solving the local optimization formulations. in all the test cases, the convergence of the po hierarchy was finite. list of symbols 0 column vector or matrix of all zeros 1 column vector or matrix of all ones a cross-sectional areas column vector asc scaled cross-sectional areas column vector ai cross-sectional area of the element i asc,i scaled cross-sectional area of the element i âr optimized cross-sectional areas at the relaxation r bi auxiliary variable associated with the element i br polynomial space basis of maximum degree r c compliance (external work) c̃ lower bound on compliance ĉ upper bound on compliance c∗ globally optimal compliance csc scaled compliance ci,i positive constant dj degree of the polynomial j e young modulus f generalized force column vector fi generalized force column vector of element i f̂ generalized force column vector associated with k̂ i element index j polynomial index ii moment of inertia of the element i i identity matrix k iteration number k stiffness matrix k̂ positive definite principal submatrix of k ki stiffness matrix of the element i l lagrangian function ` column vector of element lengths `i length of the element i m index of x mr moment matrix of the order r mr−dj localization matrix of the order r −dj ne number of elements nn number of nodes pj polynomial inequality pj polynomial matrix inequality pα,j coefficients of linear combinations of monomials r relaxation order, maximum polynomial degree s inverse of c tp parameter, web and flanges thickness t time u generalized displacement column vector v column vector of coefficients of linear combination v volume 124 vol. 26/2020 on optimum design of frame structures v volume upper bound x design variables column vector xα monomial indexed by α yβ moment associated with the basis br y column vector of moments y∗ optimal column vector of moments yr column vector of moments at r-th relaxation α vector of integers β br index η tuning parameter, 0.3 in this study ζ move limit, 0.2 in this study ε small positive number, 10−6 in this study λ lagrange multiplier column vector µ lagrange multiplier ν lagrange multiplier column vector acknowledgements the work of marek tyburec was supported by the grant agency of the ctu in prague, sgs19/033/ohk1/1t/11. martin kružík and jan zeman acknowledge the financial support by the european regional development fund through the project cz.02.1.01/0.0/0.0/16-019/0000778. references [1] m. p. bendsøe, o. sigmund. topology optimization: theory, methods, and applications. springer berlin heidelberg, berlin, heidelberg, 2003. doi:10.1007/978-3-662-05086-6. [2] w. s. dorn. automatic design of optimal structures. journal de mecanique 3:25–52, 1964. [3] a. ben-tal, a. nemirovski. lectures on modern convex optimization. society for industrial and applied mathematics, 2001. doi:10.1137/1.9780898718829. [4] g. steven, o. querin, m. xie. evolutionary structural optimisation (eso) for combined topology and size optimisation of discrete structures. computer methods in applied mechanics and engineering 188(4):743–754, 2000. doi:10.1016/s0045-7825(99)00359-x. [5] g. i. n. rozvany. structural design via optimality criteria, vol. 8 of mechanics of elastic and inelastic solids. springer netherlands, dordrecht, 1989. doi:10.1007/978-94-009-1161-1. [6] j. fiala, m. kočvara, m. stingl. penlab: a matlab solver for nonlinear semidefinite optimization. arxiv preprint arxiv:13115240 2013. 1311.5240. [7] j. b. lasserre. an introduction to polynomial and semialgebraic optimization. cambridge university press, cambridge, 2015. doi:10.1017/cbo9781107447226. [8] mosek aps. the mosek optimization toolbox for matlab manual. version 8.1., 2017. [9] j. gallier. geometric methods and applications. springer new york, 2011. doi:10.1007/978-1-4419-9961-0. [10] m. kočvara, m. stingl. pennon: a code for convex nonlinear and semidefinite programming. optimization methods and software 18(3):317–333, 2003. doi:10.1080/1055678031000098773. [11] j. b. lasserre. global optimization with polynomials and the problem of moments. siam journal on optimization 11(3):796–817, 2001. doi:10.1137/s1052623400366802. [12] d. henrion, j. b. lasserre. convergent relaxations of polynomial matrix inequalities and static output feedback. ieee transactions on automatic control 51(2):192–202, 2006. doi:10.1109/tac.2005.863494. [13] d. henrion, j. b. lasserre, j. löfberg. gloptipoly 3: moments, optimization and semidefinite programming. optimization methods and software 24(4-5):761–779, 2009. doi:10.1080/10556780802699201. [14] j. löfberg. yalmip : a toolbox for modeling and optimization in matlab. in in proceedings of the cacsd conference. taipei, taiwan, 2004. [15] h. e. curto, l. a. fialkow. the truncated complex k-moment problem. transactions of the american mathematical society 2000. doi:10.1090/s0002-9947-00-02472-7. [16] d. henrion, j. b. lasserre. detecting global optimality and extracting solutions in gloptipoly. lecture notes in control and information sciences 312:293–310, 2005. doi:10.1007/10997703_15. 125 https://doi.org/10.1007/978-3-662-05086-6 https://doi.org/10.1137/1.9780898718829 https://doi.org/10.1016/s0045-7825(99)00359-x https://doi.org/10.1007/978-94-009-1161-1 1311.5240 https://doi.org/10.1017/cbo9781107447226 https://doi.org/10.1007/978-1-4419-9961-0 https://doi.org/10.1080/1055678031000098773 https://doi.org/10.1137/s1052623400366802 https://doi.org/10.1109/tac.2005.863494 https://doi.org/10.1080/10556780802699201 https://doi.org/10.1090/s0002-9947-00-02472-7 https://doi.org/10.1007/10997703_15 acta polytechnica ctu proceedings 26:118–126, 2020 1 introduction 2 solution techniques to frame optimization 2.1 problem statement 2.2 optimality criteria 2.3 nonlinear semidefinite programming 2.4 polynomial optimization 2.4.1 solution process 2.4.2 recognizing global optimality 3 sample problems 3.1 cantilever beam 3.2 10-beam frame structure 3.3 i-shaped girder with self-weight 4 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0094 acta polytechnica ctu proceedings 22:94–98, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app using secondary raw materials in lightweight open-structure concrete with good utility properties martin sedlmajer∗, jiří zach, jan bubeník brno university of technology, faculty of civil engineering, centre admas, purkyňova 139, 612 00 brno, czech republic ∗ corresponding author: sedlmajer.m@fce.vutbr.cz abstract. the paper presents the results of research in lightweight concrete with open structure made using a lightweight porous foam-glass aggregate produced from recycled glass powder. the goal was to develop lightweight concrete. in order to achieve the best possible properties while reducing binder content, the concrete was reinforced with by-product fibres, which helped reduce the weight of the concrete while delivering satisfactory mechanical properties. in the paper are proposed lightweight concrete with open structure made using foam-glass aggregate. mechanical, thermal-insulating and acoustic properties were determined on lightweight concrete. designed concrete is only made of crushed lightweight foam-glass aggregate with a combination of portland cement with the option of adding recycled pet fibres. the new concretes possess a very good ratio of thermal insulation to mechanical properties as well as good sound absorption. keywords: lightweight concrete, lightweight aggregate, recycled fibres, thermal insulation. 1. introduction concrete is nowadays one of the most widespread building materials and is indispensable for modern civil engineering. as the industry develops, hand in hand with the performance requirements for building materials, lightweight concrete has assumed a position where it cannot be replaced by any other kind of concrete. lightweight concrete combines thermal insulation and acoustic properties with good mechanical properties and thus can be used in the construction of new types of single-layer structures with multifunctional properties. the reduced bulk density of this concrete saves material and energy during construction. depending on the requirements, there are several methods of manufacturing lightweight concrete. the most common means of weight reduction is using a lightweight filler instead of a dense aggregate. this filler is usually a lightweight aggregate. it can be natural or artificially manufactured. artificial aggregates are used especially in cases where no sources of natural aggregate are available. one such material is foam glass, which is made from recycled glass powder. foam glass is manufactured artificially and is a very good source of lightweight aggregate. there are several varieties of lightweight foam glass aggregate, depending on manufacturing technology. they can be round pellets or sharp, angular particles produced by crushing larger sheets of foam glass. both methods yield different fractions, which are then separated into broad or narrow according to need. this is important for optimal formula design, specifically, the correct particle size distribution. foam-glass aggregate has properties that make it suitable for use in both reinforced and non-reinforced applications, specifically fills, foundation insulation, or concrete aggregate [1], where it provides a considerable weight reduction and alters both freshand hardenedstate properties. the aggregate reduces bulk density and thus improves thermal insulation properties far above traditional concrete that uses natural dense aggregate [2]. a disadvantage of foam-glass aggregate are typically poorer mechanical properties, resulting from the reduced bulk density because aggregate is a dominant component of every concrete and the main provider of mechanical properties. lightweight concretes with foam-glass aggregate can also be used in places where ordinary dense concrete is unsuitable, e.g. lightweight structures, thermal insulation and eliminating thermal bridges [3], improving fire resistance [4], etc. it still preserves all of the benefits of concrete, the most important being its mouldability. 2. test mixtures and methodology the research was performed with a lightweight crushed foam-glass aggregate produced by refaglass s.r.o. it is an angular aggregate made by crushing larger pieces of foam glass and separating them by desired fraction. the feedstock for foam glass manufacturing is recycled glass powder, most often types of bottle glass that are difficult to use elsewhere in the glass industry. in addition to glass powder, another important ingredient for foam glass manufacturing is the foaming agent, which causes expansion at high temperatures (around 900 °c) in electric furnaces. the refaglass aggregate 94 http://dx.doi.org/10.14311/app.2019.22.0094 http://ojs.cvut.cz/ojs/index.php/app vol. 22/2019 using secondary raw materials in lightweight open-structure concrete. . . component/mixture no. 1 2 3 4 5 6 7 portland cement [kg] 250 200 150 130 100 130 100 aggregates from foam glass 4–16 mm [l] 1000 1000 1000 1000 1000 1000 1000 polycarboxylate ether superplasticizer [kg] 2 1.6 1.2 1 0.8 1 0.8 water 106 61 57 53 51 53 65 pet fibres [kg] 6.5 5.0 table 1. mixtures of lightweight concrete. used in this research was of a 4–16 mm fraction (particle size of 4 to 16 mm). this fraction was used in the manufacturing of the experimental lightweight concretes with open structure. drawing on previous research, the mixtures of the new concretes were designed. there were 7 mixtures in total, out of which mixtures 1-5 contained no fibres and differed in binder content, which descended from 250 kg/m3 down to 100 kg/m3. mixtures 6 and 7 contained recycled pet fibres provided by ciur a.s. at the amount of 5% of cement weight. plasticiser content was uniform in all mixtures as well; 0.8% of cement weight. table 1 shows the composition of the mixtures. the lightweight concretes were observed for key properties so as to compare their mixtures in fresh and hardened state. the basic properties to be observed included freshand hardened-state bulk density, compressive strength, and thermal conductivity. mixtures were produced according to the formulae described above. the following tests were performed on the fresh mixtures: • determination of consistency in fresh concrete en 12350-2 [5], • determination of bulk density in fresh state en 12350-6 [6]. and in hardened state: • determination of bulk density in hardened state en 12390-7 [7], • determination of compressive strength en 12390-3 [8], • determination of thermal conductivity coefficient en 12667 [9] (according to iso 8301), • determination of sound absorption coefficient iso 10534-1 [10]. the recycled pet fibres had the thickness of ca. 13µm and length of 6.3 mm. figure 1 shows a recycled pet fibre. 3. results and discussion the new lightweight concretes with open structure were first tested for consistency [5] and bulk density in fresh state [6]; the values are listed in table 2 and in figure 2. it was necessary to watch the consistency because of the danger of cement paste segregation due to the great difference between the bulk density of both figure 1. detail image of pet fibre. materials. this is why the cement paste consistency was set so as to ensure that all the foam-glass particles are completely coated in the paste without it flowing through and accumulating at the bottom. once correct consistency was determined, attention was paid to measuring hardened state properties, primarily compressive strength [8] and bulk density [7]. table 3 shows the measured values. figure 3 shows the values of compressive strength and bulk density in a hardened state. the reduction of binder content, which entailed reducing cement content from the original 250 kg/m3 (mixture 1) down to 100 kg/m3 (mixture 5), which is more than a half, resulted in a loss of mechanical properties, making the concrete poorly bonded and prone to crumbling. looking at mixture 7, the combination of 100 kg of portland cement with recycled pet fibres improved the concrete’s bond and compressive strength, which reached 1.2 mpa. the benefit of the fibres is clear in this case. a similar effect can be seen in mixtures 4 and 6, both of which contained pet fibres and thus profited from an improvement in mechanical properties, specifically compressive strength, by more than 100%. next, the hardened concrete specimens were tested for steady-state thermal conductivity by the hot plate method (at a mean temperature of +10 °c and a temperature gradient of 10 °c) [9]. before that the test the specimens were dried at 105 °c down to stable mass. the results are listed in table 4 and plotted in figure 4. 95 m. sedlmajer, j. zach, j. bubeník acta polytechnica ctu proceedings properties/mixture no. 1 2 3 4 5 6 7 bulk density in fresh state [kg] 860 740 580 500 510 670 540 slump test [mm] 20 0 0 0 0 0 0 table 2. properties of the lightweight concrete in fresh state. figure 2. bulk density and consistency of fresh lightweight concrete. properties/mixture no. 1 2 3 4 5 6 7 bulk density in hardened state [kg] 780 670 550 460 500 600 490 compressive strength [mpa] 3.6 2.9 1.5 0.8 0 1.8 1.2 table 3. bulk density and compressive strength of lightweight concrete in hardened state. figure 3. compressive strength and bulk density of hardened lightweight concrete. properties/mixture no. 1 2 3 4 5 6 7 thermal conductivity coefficient [w/(m·k)] 0.1529 0.1419 0.1223 0.1112 0 0.1214 0.1013 table 4. thermal conductivity coefficient of lightweight concrete in hardened state. 96 vol. 22/2019 using secondary raw materials in lightweight open-structure concrete. . . figure 4. thermal conductivity coefficient and bulk density of hardened lightweight concrete. figure 5. sound absorption coefficient of hardened lightweight concrete 2. figure 6. overview of the calculated single-number values of sound absorption čsn en 1793. 97 m. sedlmajer, j. zach, j. bubeník acta polytechnica ctu proceedings the results of thermal conductivity tests do not show any major effect of the fibres on thermal conductivity. as could be expected, however, what matters more is cement content. the last parameter was the sound absorption coefficient, which expresses how well a material can absorb incoming acoustic energy. the measurements were performed using an acoustic interferometer according to iso 10534-1 [10]. the measurement was made within 100–6300 hz in one-third octave bands. the measured values were used in the calculation of single-number values of sound absorption dlα [db] in accordance with results are listed in figure 6. the progress of the sound absorption coefficient is plotted in figure 5. the results show that mixtures 2 and 6 perform the best. 4. conclusions lightweight crushed foam-glass aggregate can be used for manufacturing open-structure concrete with a good ratio of thermal insulation and mechanical properties. the results show that a reduced amount of cement results in poorer mechanical properties, but better thermal insulation. reducing the cement content down to 100 kg/m3 resulted in a weak, insufficiently bonded concrete. the addition of recycled pet fibres to the mixture with 130 kg of cement raised its compressive strength by 125% from 0.8 mpa to 1.8 mpa. mixture 7 with 100 kg of cement and pet fibres had compressive strength of 1.2 mpa. this corresponds to a mixture without fibres with a cement content of ca. 140 kg of cement. it can thus be said that the addition of pet fibres brought a dramatic improvement of the concrete’s properties and enabled a significant reduction of cement content without damaging the mechanical properties. in mixture 7 the reduction of cement content and addition of fibres further reduced the concrete’s thermal conductivity. its final value was 0.1013 w/(m·k), which makes this concrete suitable for constructing structures with excellent thermal insulation properties. the newly developed lightweight concrete also shows very good acoustic properties (compared to ordinary dense concrete); mixtures 2 and 6 performed the best in this regard. acknowledgements this paper was written under the project no. lo1408 “admas up advanced materials, structures and technologies”, supported by the ministry of education, youth and sports under the “national sustainability programme i” and project fv20086 “development of light new-age building materials with utilization of light-weight aggregate based on waste glass powder” supported by the ministry of industry and trade. references [1] m. limbachiya, m. s. meddah, s. fotiadou. performance of granulated foam glass concrete. construction and building materials 28(1):759–768, 2012. doi:10.1016/j.conbuildmat.2011.10.052. [2] m. b. janetti, t. plaz, f. ochs, et al. thermal conductivity of foam glass gravels: a comparison between experimental data and numerical results. energy procedia 78:3258–3263, 2015. doi:10.1016/j.egypro.2015.11.713. [3] s. real, m. g. gomes, a. m. rodrigues, j. a. bogas. contribution of structural lightweight aggregate concrete to the reduction of thermal bridging effect in buildings. construction and building materials 121:460– 470, 2016. doi:10.1016/j.conbuildmat.2016.06.018. [4] c.-g. go, j.-r. tang, j.-h. chi, et al. fire-resistance property of reinforced lightweight aggregate concrete wall. construction and building materials 30:725–733, 2012. doi:10.1016/j.conbuildmat.2011.12.081. [5] european standard en 12350-2 testing fresh concrete part 2: slump-test. european standard, 2009. [6] en 12350-6 testing fresh concrete part 6: density. european standard, 2009. [7] european standard en 12390-7 testing hardened concrete part 7: density of hardened concrete. european standard, 2009. [8] european standard en 12390-3 testing hardened concrete part 3: compressive strength of test specimens. european standard, 2009. [9] european standard en 12667 thermal performance of building materials and products determination of thermal resistance by means of guarded hot plate and heat flow meter methods products of high and medium thermal resistance. european standard, 2001. [10] iso 10534-1 acoustics determination of sound absorption coefficient and impedance in impedance tubes part 1: method using standing wave ratio. european standard, 1999. 98 http://dx.doi.org/10.1016/j.conbuildmat.2011.10.052 http://dx.doi.org/10.1016/j.egypro.2015.11.713 http://dx.doi.org/10.1016/j.conbuildmat.2016.06.018 http://dx.doi.org/10.1016/j.conbuildmat.2011.12.081 acta polytechnica ctu proceedings 22:94–98, 2019 1 introduction 2 test mixtures and methodology 3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2021.31.0005 acta polytechnica ctu proceedings 31:5–9, 2021 © 2021 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague high-speed rail network and periodic timetable: a comparative analysis of operational concepts michal drábek∗, dominik mazel, jiří pospíšil czech technical university in prague, faculty of transportation sciences, department of logistics and management of transport, horská 3, 128 03 prague 2, czech republic ∗ corresponding author: xdrabek@fd.cvut.cz abstract. this paper compares chosen european high-speed railway (hs) networks in terms of their offer of hs passenger service. the criteria chosen for comparison are network topology, degree of service periodicity and degree of coordination between subsequent services. only services with hs trains are taken into account. as a result, each examined network is classified according to prevailing approach to passenger service – either line/service (ls) approach, where transfer connections are in general not anticipated, or network (n) approach, with regular (mostly periodic) public transport lines and periodic transfer connections between them. the comparative analysis has shown that geography had crucial impact not only on national (or regional) hs line network, but on the hs operational concept as well. on trunk hs lines, which connect most populated agglomerations in particular country, there is always – at least during peak times – some form of periodic service, despite compulsory seat reservation (except state-owned carriers in austria and germany). half of analyzed networks can be characterized by n approach – at least on trunk hs lines or within central "core" part of hs network. for czech hs network, authors recommend to define a core network with application of integrated periodic timetable. keywords: high-speed railway, operational concept, passenger service, periodicity. 1. introduction high-speed (hs) railway service network is a specific form of long-distance railway network. the reasons are longer average trip distance (due to higher commercial speed) and higher planned load factor of hs trains (due to rolling stock and infrastructure cost). hs railway very often beats car even in door-to-door travel time, which leads to unprecedented modal shift towards railway. for estimation of modal-split based on comparation of travel time, see for instance vávra and janoš [1]. either this network is more complex – mostly in the case of state-owned carrier(s), or more simple. two approaches for design of hs operational concept can be distinguished: • line/service (ls) approach • network (n) approach ls approach focuses on origin-destination (od) pairs and times with the highest demand. these od pairs are mostly radial, i.e. to/from central agglomeration. transit trips through this centre are perceived as secondary, i.e. there can be a connection in the central station, but transfer time can be very high, as timetable was not adjusted for this requirement. n approach is based on services which operate in regular public transport lines. all services run through (at least mostly) equal route, with some degree of periodicity (either more service pairs daily, or with some period – at least some services, typically during peak times in relevant direction). periodic transfer connections between these public transport lines are anticipated. the extremely periodic concept is integrated periodic timetable (ipt, taktfahrplan). for more detailed explanation, see for instance weidmann [2]. the aim of this paper is to compare chosen national hs networks in terms of their topology and degree of periodicity of hs passenger services. finally, each network operation will be assigned to (predominant) ls or n approach. 2. methodology and scope of the analysis data on european hs railway network were taken from uic high speed rail 2018 publication [3]. the timetable data, as of 2019/2020 railway timetable, were gathered from timetable websites or online publications of particular carriers [4–14]. for each analyzed national or regional network, following criteria were chosen: • network topology . one or two principal hs lines . star (one center and radial hs lines) 5 https://doi.org/10.14311/app.2021.31.0005 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en m. drábek, d. mazel, j. pospíšil acta polytechnica ctu proceedings . raster network • prevailing degree of periodicity . no periodic service . periodic service in peaks only . partially periodic service on trunk routes . almost fully periodic – with some minor irregularities . fully periodic (with possible additional peak service) following hs networks were analyzed: • austria (öbb and westbahn) • germany (ice) • thalys (belgium, netherlands, france and germany) • france (tgv incl. oui) • italy (fs – frecce trains and .italo) • spain 3. results of the analysis 3.1. austria two principal hs corridors are planned in austria. the first one, western railway (westbahn) is partially in operation – from vienna to linz, with a short section westwards. in this section, there is mostly a parallel hs line to a conventional line, with frequent connecting junctions. further sections are in construction or planned – new lower inn valley railway to innsbruck and brenner base tunnel to italy. the next corridor, currently in construction, goes southwestwards from vienna through graz and klagenfurt. despite axis nature of these hs corridors, the operation there is fully integrated into national integrated periodic timetable [4]. on the only austrian hs line in operation, there are three hs carriers. all operate in zero-symmetric periodic timetables. state-owned öbb-personenverkehr ag operates railjet network, with regular periodic service. the first railjet axis, prague – brno – vienna – graz, is operated in 2-hour period. there are periodic transfer connections with railjet to/from western railway. however, this public transport line does not currently use any hs infrastructure [5]. on the western railway, there operate even two segments of hourly raijlet services between vienna and salzburg [6]. the lower segment fulfills function of intercity service, but uses the hs line from vienna to tullnerfeld junction only. the higher segment, named railjet xpress, serves only vienna, st. pölten, linz and salzburg. from vienna, both segments depart interposed in common 30-min period, but in salzburg they meet each other around minute zero. so, periodic transfer connections are enabled. each 2-hours, an unit from railjet xpress proceeds directly to munich. from salzburg westwards, there is hourly intercitylike service, mostly operated by railjet xpress, but occasionally by railjet, eurocity etc. similar kind of irregularities occurs from budapest to vienna, where mostly railjet xpress services operate, proceeding directly westwards. the service period from/to budapest varies from 30 min to 2 hours. second hs carrier is german state-owned deutsche bahn (db), with intercity-express (ice) services to frankfurt, mostly with 2-hour service period. last hs carrier is (mostly) privately-owned company westbahn (not to be confused with western railway), which operates in mostly 2-hour period between vienna and salzburg [7]. austrian hs operation can be classified as periodic, with few irregularities, and n approach. 3.2. germany german hs railway network creates a raster, in accordance with polycentric settlement of the country. on the main (trunk) routes, there is mostly periodic offer of ice services. basic service period of a long-distance public transport line is usually 2 hours. there are many minor deviations in departure or arrival minutes, or in presence of intermediate stops of secondary significance (i.e. cities with population circa 100,000). another deviations occur in days of service operation. for some services, there are alternative destinations, which may result in a gap in periodic service during some part of the line. there are some cases of interpositions into half period on trunk routes (and even in the case of alternative routes, but same od-pair, e.g. nuremberg – hamburg via hannover or via erfurt). on particular sections between major cities (which are usually principal railway nodes as well), one line of hs trains serves as "lower segment", serving middle-sized cities. other lines of hs trains run through without stop [8]. german hs operation can be considered predominantly periodic, with n approach, albeit with many minor irregularities. 3.3. thalys thalys is a joint venture of french sncf (majority owner) and belgian nmbs/sncb state-owned rail carriers. although predominantly periodic, thalys services are not interconnected at all. the reason is probably comparably short travel distances, so any transfer would disturb the passengers. the service period is mostly 1 or 2 hours (peak), between paris and brussels even 30 min during peaks. from brussels, there are 3 hs lines: to paris nord station, to cologne and to amsterdam. from amsterdam, 13 pairs of services run to brussels, which include 10 pairs that proceed to paris. the period is 1 hour, but there is a 3 hours peak gap from amsterdam. from cologne, 4 pairs of services run to brussels, but there are more additional german ice hs services that make together 1 hour period (with 2 pairs of services left out). from paris, there is 30 min service period during both morning and afternoon 6 vol. 31/2021 high-speed rail network and periodic timetable peaks, but there is no 1 hour period kept all day long. peaks, there are 1.5 hour gaps between subsequent services [9]. thus, despite periodic services (with many irregularities), thalys network is characterized with ls approach. 3.4. france french geography is characterized by strong dominance of île-de-france (paris agglomeration), where one out of six french inhabitants lives, and by transport network converging to paris. hs network is no exception. from four paris dead-end stations, a four-pointed star network (as of 2020) runs out. farther away, some hs railway lines (lgv) branch to various destinations. this network is completed by a hs bypass line, lying eastwards from paris, with stops at roissy – charles de gaulle airport and marne-la-vallée-chessy. french hs service offer is mainly focused on direct connection of paris with as many cities as possible – for smaller cities, there can be circa 2 to 4 pairs of train à grande vitesse (tgv) services per day. this policy leads to more complex branching of tgv routes – much beyond branching of hs railway lines. sometimes there run two coupled tgv units from paris, and from some node station they proceed to different destinations. by nature, such policy makes periodic service offer hardly possible. moreover, even on trunk sections of hs lines, common for more end destinations (e.g. to lyon or to strasbourg), there is almost no periodicity of service – only some groups of subsequent services during afternoon peak from paris (or during morning peak to paris). there are two more factors that increase complexity and ambiguity of tgv service concept. the first one is phenomenon of stations, located directly on a hs line. in some cases, such station lies "in the fields", located between more comparably populated cities. otherwise, this station serves a large city, but lies few to circa 20 km remote. so, there is another station in the city centre. in most such cases, there are both tgv services which serve the station on a hs line, and the others, which run to the centre. the second factor is phenomenon of "tgv intersecteurs" services, which stop on bypass lgv, but not in paris itself [10, 11]. french hs operations can be considered as practically not periodic, with ls approach. 3.5. italy italian hs network consists of three north-south corridors, which merge together in bologna. then, a single hs line (including direttissima hs line, whose construction started as first one in europe) proceeds via rome to naples and salerno. from naples, there is a mostly completed hs branch to adriatic coast (to foggia and bari). moreover, in sicily, there are two partially complete coastal hs lines from messina to palermo and to catania. at present, they are served by conventional trains only. in northern italy, there is an east-west corridor from venice to turin. on the main (trunk) routes, there is mostly periodic offer of frecce (literally "arrows") hs trains, but, contrary to germany, these trains do not always operate in regular public transport lines. the service period of these trains, operated by state-owned fs (trenitalia) carrier is 1 or 2 hours. the frecce offer concept consists of three products, which differ in type of trainset: • red arrow ("frecciarossa") stands for hs train with maximum speed 300 km/h • silver arrow ("frecciargento") stands for tilting hs trains with maximum speed 250 km/h • white arrow ("frecciabianca") stands for classical trainsets with maximum speed 200 km/h since majority of italian hs lines are designed for 300 km/h, there are obviously services that cannot use this maximum speed, because of considerably lower maximum trainset speed. according to actual timetable [12] this situation occurs few times a day, and there is no lengthening of regular running times. in general, such heterogeneity of train paths leads to substantial drop in railway capacity utilization – see for instance janoš and kříž [15]. the competitor – nuovo trasporto viaggiatori (ntv), with services labeled as ".italo" – operates public transport lines between milan and venice, and on three principal hs lines, further together to rome and naples, with few pairs of services proceeding to salerno. single pairs of services proceed directly to further chosen cities in northern italy. .italo services depart in general at regular, periodic times. there are five public transport lines (milan – venice, 3 hs corridors), with few lengthenings northwards. from milan to rome, besides line turin – rome (– naples) there is in addition a non-stop line between these two cities. both lines operate hourly, with few service pairs left out. another public transport lines operate in periods varying from 1 to 3 hours during the day, sometimes with even longer off-peak gap. transfer connections are announced, but not anticipated in planning – the transfer time can reach up to 1.5 hours [13]. as a conclusion, hs passenger service in italy (both state and private carriers) can be classified as predominantly periodic (at least on trunk routes), with n approach (state carrier fs), since there are periodic transfers to non-hs services with transfer time of 30 min or less. 3.6. spain spanish hs railway network follows the country’s specific – mostly mountainous – geography. besides madrid agglomeration and few middle-sized agglomerations, vast majority of population lives in coastal areas or in major valleys, where most of the largest 7 m. drábek, d. mazel, j. pospíšil acta polytechnica ctu proceedings spanish cities are located. contrary to conventional network with wide (iberian) gauge, hs network is constructed with standard gauge. the hs railway network is star-shaped, with two central stations in madrid. chamartín station is a terminus of hs lines to northern and western coast. atocha station is a terminus of remaining hs network (eastern and southern coast and some inland destinations like toledo and granada). there are two main segments of hs service – ave, and avant with alvia [14]. ave is a distinctly hs service with operating speed up to 300 km/h. avant and alvia are lower hs segments with operating speed up to 250 km/h. while avant serves middle distances on hs lines, alvia trains are equipped with gauge-changing device (for some services, talgo trainsets are deployed). thus, alvia trains offer direct services far beyond hs network (mostly to regional centres in coast areas – mainly northwards from madrid). since many destinations and lower frequency (few service pairs daily), there is no regular period of alvia services. in some cases, two alvia trains (or two other hs trains) run coupled, at least during part of their route. spanish hs network is, in general, characterized by non-periodic service, and transfer connections are not anticipated. however, on trunk routes (from madrid to barcelona, sevilla, malaga and valencia) and for some shorter distances (e.g. madrid – toledo), there is periodic service at least during peak times. the period is mostly 1 hour, with almost all-day 30 min period between madrid and barcelona, with additional peak services. there are even few direct hs services between barcelona and sevilla, and euromed coastal hs services from barcelona to valencia. the periodic service is not without various irregularities – either different stopping pattern for each service, or some 1.5 hour gaps between services. so, hs trains to toledo depart from madrid hourly, mostly in the minute 50, but in off-peak time in the minute 20. as a conclusion, spanish hs passenger service can be characterized as partially periodic, with ls approach. 3.7. structured comparison of analyzed networks structured comparison of analyzed networks and hs operational concepts is displayed in table 1. 4. conclusions and recommendations the comparative analysis has shown that geography had crucial impact not only on national (or regional) hs line network, but on the hs operational concept as well. if there is a clearly dominant central agglomeration, the hs passenger service is usually driven by ls approach. the only exception is austria. on trunk hs lines, which connect most populated agglomerations in particular country, there is always at least during peak times some form of periodic service, despite compulsory seat reservation (except state-owned carriers in austria and germany). half of analyzed networks can be characterized by n approach at least on trunk hs lines or within central "core" part of hs network. on the other hand, because hs trains are very expensive, the carriers (even state-owned ones) focus on operational efficiency, at the expense of timetable simplicity and regularity. a notorious example is a 1.5-hour gap between subsequent services. czech republic stands before strategic decisions on high-speed operational concept, since planning of hs lines is already almost in the beginning. the authors recommend to define core part of hs network with application of integrated periodic timetable, and operation without compulsory seat reservation for at least part of a train. open access services should be enabled in such way that they would not disturb this concept. acknowledgements this paper is a partial output of research project ck01000004 "efficient operational concept for rapid services", which is co-financed from the state budget by the technology agency of the czech republic under the transportation 2020+ (doprava 2020+) programme – http://www.tacr.cz, and by ctu in prague, faculty of transportation sciences. the authors hereby acknowledge both sponsors for funding for this project. references [1] r. vávra, v. janoš. analysis of competitiveness of railway passenger transport between rakovník and prague, 2017. acta polytechnica ctu proceedings 11:86–90, https://doi.org/10.14311/app.2017.11.0086. [2] u. weidmann. system-und netzplanung. band 1.1. system-und netzplanung des personenverkehrs, may 2008. ivt eth zürich, zürich. [3] uic. uic high speed rail 2018. [2020-03-22], https://uic.org/img/pdf/ic_high_speed_2018_ ph08_web.pdf. [4] öbb-infrastruktur ag. schienennetznutzungsbedingungen 2020 der öbb-infrastruktur ag. [2020-03-22], https://infrastruktur.oebb.at/de/ geschaeftspartner/schienennetz/snnb/snnb-2020/ schienennetz-nutzungsbedingungen-2020.pdf. [5] české dráhy. jízdní řád 2020, dálkové vlaky, ex3, r9, r19. [2020-03-22], https://www.cd.cz/images/cdosn/ trains/ex3-r9-r19-2020.pdf. [6] öbb kundenservice. schnell und komfortabel richtung westen. [2020-03-22], https://www.oebb.at/dam/jcr: 21c69cd4-78f6-472e-a907-637f9ee11a08/ folder-go-west.pdf. 8 http://www.tacr.cz https://doi.org/10.14311/app.2017.11.0086 https://uic.org/img/pdf/ic_high_speed_2018_ph08_web.pdf. https://uic.org/img/pdf/ic_high_speed_2018_ph08_web.pdf. https://infrastruktur.oebb.at/de/geschaeftspartner/schienennetz/snnb/snnb-2020/schienennetz-nutzungsbedingungen-2020.pdf https://infrastruktur.oebb.at/de/geschaeftspartner/schienennetz/snnb/snnb-2020/schienennetz-nutzungsbedingungen-2020.pdf https://infrastruktur.oebb.at/de/geschaeftspartner/schienennetz/snnb/snnb-2020/schienennetz-nutzungsbedingungen-2020.pdf https://www.cd.cz/images/cdosn/trains/ex3-r9-r19-2020.pdf https://www.cd.cz/images/cdosn/trains/ex3-r9-r19-2020.pdf https://www.oebb.at/dam/jcr:21c69cd4-78f6-472e-a907-637f9ee11a08/folder-go-west.pdf https://www.oebb.at/dam/jcr:21c69cd4-78f6-472e-a907-637f9ee11a08/folder-go-west.pdf https://www.oebb.at/dam/jcr:21c69cd4-78f6-472e-a907-637f9ee11a08/folder-go-west.pdf vol. 31/2021 high-speed rail network and periodic timetable hs network topology degree of periodicity hs oper. approach austria 1 principal route fully periodic n germany raster almost fully periodic n thalys star periodic ls france star + branching periodic in peaks only ls italy raster + 1 principal route + branching periodic (trunk routes) n spain star+branching+raster part. periodic (trunk routes) ls table 1. structured comparison of networks and hs operational concepts. [7] westbahn management gmbh. where do you want to travel? [2020-03-22], http://westbahn.at. [8] db vertrieb gmbh. die bahn. timetable & booking. [2020-03-22], http://www.bahn.com. [9] thalys. book your next trip. [2020-03-22], http://www.thalys.com. [10] sncf. itinerary. [2020-03-22], http://www.sncf.com. [11] oui.sncf. votre fiche horaire. [2020-03-22], https: //www.oui.sncf/services-train/fiches-horaires. [12] trenitalia. gruppo ferrovie dello stato italiane. f. dello stato italiane. [2020-03-22], http://www.trenitalia.com. [13] .italo. nuovo trasporto viaggiatori. where are you going? [2020-03-22], http://www.italotrento.it. [14] renfe. [2020-03-22], http://www.renfe.com. [15] v. janoš, m. kříž. infrastructure parameters affecting capacity of railways in ten-t, 2016. acta polytechnica ctu proceedings 5:22–25, https://doi.org/10.14311/app.2016.5.0022. 9 http://westbahn.at http://www.bahn.com http://www.thalys.com http://www.sncf.com https://www.oui.sncf/services-train/fiches-horaires https://www.oui.sncf/services-train/fiches-horaires http://www.trenitalia.com http://www.italotrento.it http://www.renfe.com https://doi.org/10.14311/app.2016.5.0022 acta polytechnica ctu proceedings 31:5–9, 2021 1 introduction 2 methodology and scope of the analysis 3 results of the analysis 3.1 austria 3.2 germany 3.3 thalys 3.4 france 3.5 italy 3.6 spain 3.7 structured comparison of analyzed networks 4 conclusions and recommendations acknowledgements references doi:10.14311/app.2021.30.0012 acta polytechnica ctu proceedings 30:12–17, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app on comparison of 3d isogeometric timoshenko and bernoulli beam formulations edita dvořáková∗, bořek patzák czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: edita.dvorakova@fsv.cvut.cz abstract. application of isogeometric analysis (iga) for curved beams is very convenient for its ability of exact representation of curved geometries. several beam formulation has been presented since the introduction of iga. in this paper, two different beam formulations are presented: bernoulli beam formulation of a. m. bauer et al. [1], and timoshenko beam element introduced by g. zhang et al. [2]. both beam elements are implemented and their performance is documented on the fully threedimensional example of helicoidal spring. keywords: bernoulli theory, curved beam element, isogeometric analysis, nurbs, timoshenko theory. 1. introduction isogeometric analysis is an alternative of standard finite element analysis introduced in 2005 by t. hughes [3]. the crucial impulse for its development was given by the desire for the automatic connection between computer-aided design (cad) and finite element analysis. this goal has been achieved by the application of the same nurbs-based geometry description in cad also for the unknown approximation in the analysis. the use of cad basis functions enables the exact description of arbitrarily curved geometries including curved beams. the focus of this paper is placed on the isogeometric analysis of spatial curved beams. two different beam formulations are presented and their performance is compared on a benchmark problem. the first beam formulation is based on bernoulli beam theory, while the timoshenko kinematics is assumed in the second beam element presented. 2. nurbs-based analysis non-uniform rational b-splines (nurbs) [4] are the most widespread technology used in cad industry. their fitness for cad is rooted in the ability of the exact description of a vast range of geometries, including the conic sections widely used in architectural design. there are many effective algorithms developed for handling nurbs objects which can be also very useful for isogeometric analysis. the computational domain in isogeometric analysis (iga) is firstly divided into patches, which are further divided into knotspans (often referred to as elements in iga), see fig. 1. the nurbs curve is obtained by linear combination of cartesian coordinates of the control points p and nurbs functions rpi c(ξ) = n! i=1 r p i (ξ)pi. (1) nurbs functions are rational polynomials which are generated by weighting b-spline functions with a weights associated with a control points. b-spline functions are piecewise polynomials and can be derived recursively starting with a piecewise constant functions, see [4] for the detailed description. the use of nurbs basis introduces some special aspects to the analysis. the nurbs basis functions are generally non-iterpolatory, thus the degrees of freedom in the control points lack the direct physical meaning. additionally, the higher continuity between knotspans can be achieved naturally using nurbs, unlike c 0 in standard fem. one of the main features of iga is the fact, that the basis can be enriched without altering geometry representation using knot insertion and degree elevation algorithms. 3. bernoulli beam element formulation the first presented beam formulation is adopted from the work of a. m. bauer et al. [1]. the element is based on bernoulli beam theory, thus the orthogonality of the cross-section to the center line after deformation is assumed to be preserved and the crosssectional dimensions remain unchanged. the formulation accounts for the geometrically nonlinear behavior, nevertheless only the linear behavior is studied in the presented examples. the element has four degrees of freedom in each control point: three global displacements u, v, and w and an additional degree of freedom corresponding to the rotation around the centerline ψ. 12 http://dx.doi.org/10.14311/app.2021.30.0012 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 comparison of timoshenko and bernoulli beams n4,2 ⇠ knotspan patch continuity: c0 c0c1c1 c1 knotspan p = 2, ⌅ = {0, 0, 0, 0.5, 1, 1, 1} n1,2 n2,2 n3,2 knots control points p1 p2 p4p3 parametric space physical space x y figure 1. description of b-spline (nurbs) finite element geometry. the geometry is modeled by one patch consisting of two knotspans with quadratic nurbs approximation. in the following, the standard notation using upper-case and lower-case letters for the undeformed and deformed configuration, respectively, is adopted. the geometry description is illustrated in figure 2. the three-dimensional beam approximation can be reduced to the center line given by the position vector xc (xc) xc = ! i r p i x̂ i, (2) xc = ! i r p i x̂ i, (3) where x̂i are the control points coordinates and rpi are the corresponding basis functions. deformed position vector is given as x̂i = x̂i + ûi, (4) where ûi is a vector of global degrees of freedom (u, v, w). the cross-section orientation is described by a moving trihedral given by the base vectors ai (ai) with i ∈ {1, 2, 3}. with assumption of bernoulli kinematics, a position vector of an arbitrary point of a beam continuum given by coordinates x (x) can be expressed as x(θ1, θ2, θ3) = xc(θ1) + θ2a2(θ1) + θ3a3(θ1), (5) x(θ1, θ2, θ3) = xc(θ1) + θ2a2(θ1) + θ3a3(θ1), (6) where θi are the convective contravariant coordinates. the first components of a moving trihedral a1 and a1 are aligned with a normalized tangents t and t, respectively, and are computed as a linear combination of the control points coordinates and the derivatives of the basis functions. the remaining components aα with α ∈ {2, 3} are orthogonal to the tangent of the center line (resp. a1) and are derived from the reference trihedral, given by a0α, t0, using two key operations used for the alignment at the desired position. these operations are illustrated in figure 3. analogical procedures denoted as λ(t, t) and r̄t(ψ) are used to align the moving trihedral of the undeformed to the deformed configuration. by substituting geometric relations into the definition of green-lagrange strain tensor, the individual components of the tensor can be derived as e11 = 1 2 (a11 − a11) " #$ % " +θ2 a2,1 · a1 − a2,1 · a1" #$ % κ21 +θ3 a3,1 · a1 − a3,1 · a1" #$ % κ31 . (7) the diagonal terms eαα as well as the shear term e23 are equal to zero. this yields from the bernoulli assumptions (undeformable cross-section). for the off-diagonal terms e12, e13 e1α = 1 2 θβ (aβ,1 · aα − (aβ,1 · aα)" #$ % κβα , (8) where (α, β) ∈ {(2, 3), (3, 2)}. the relations between strains and displacements are not linear and to the purposes of the linear analysis considered in this paper need to be linearized. the energetically conjugated stress tensor to the greenlagrange strain tensor is the second piola-kirchhoff tensor. the stiffness matrix is obtained by a standard procedure [5]. the fully derived equations for the presented beam formulation can be found in [1]. 4. timoshenko beam element formulation the second beam element presented here has been proposed by g. zhang et al. [2]. the beam formu13 edita dvořáková, bořek patzák acta polytechnica ctu proceedings a1(✓ 1) a2(✓ 1) a3(✓ 1) a1(✓ 1) a3(✓ 1) a2(✓ 1) u(✓1, ✓2, ✓3) x(✓1, ✓2, ✓3) xc(✓ 1) xc(✓ 1) x y z figure 2. illustration of the bernoulli beam in its undeformed and deformed configurations. t0 t0 a02 a03 t ⇤(t0, t) ⇤(t0, t)⇤(t0, t)a 0 3 a03 ⇤(t0, t)a 0 2 t0 a02 a03 t ⇤(t0, t)a 0 3 ⇤(t0, t)a 0 2 r̄t ( ) r̄t ( ) a3 a2 figure 3. the λ(t0, t) and r̄t(ψ) operations used for the alignment of the bernoulli beam cross-section at the current position. t(s) n(s) b(s) ut ✓t ub ✓b un ✓n r(s) figure 4. local coordination system and degrees of freedom of timoshenko beam element. lation is based o the timoshenko beam theory and is derived in local coordinate system (t, n, b), where t, n, b are tangent, normal and binormal vectors. there are six independent unknowns: tangential displacement ut, normal displacement un, binormal displacement ub and rotations θt, θn, θb (see fig. 4). a strain-displacement matrix b is defined as ε = br, (9) where ε = {εm, γn, γb, χt, χn, χb}t and r = {ut, un, ub, θt, θn, θb}t . b is derived from the following relations for membrane strain εm, shear strains γn and γb, torsional strain χt, and bending strains χn and χb εm = u′t − κun, χt = θ′t − κθn, γn = κut + u′n − τ ub − θb, χn = κθt + θ′n − τ θb, (10) γb = τ un + u′b + θn, χb = τ θn + θ′b. curvature κ is given as κ = &&&& &&&& d2r(s) ds2 &&&& &&&& (11) and torsion τ τ = dn(s) ds b(s), (12) where r(s) is the position vector and s is the curvilinear coordinate measured along the centerline of the beam. conjugated generalized stress vector contains normal and shear forces, and torsional and bending moments. similarly to the previous formulation, by following standard procedure the stiffness matrix is obtained. 4.1. numerical locking the well-known drawback of the timoshenko beam element formulations is numerical locking. locking phenomena rises from the use of the same approximation order for displacements and for rotations, which leads to the shear strains of a higher order than bending moments. several methods for unlocking beam elements are available, b̄-method [6] is used for the purposes of 14 vol. 30/2021 comparison of timoshenko and bernoulli beams this study. this method is based on the projection of the membrane and shear strains onto a basis of a lower order, leading to the removal of the locking. see [6] for the detailed description of the method. 5. numerical examples both presented beam formulations have been implemented and their performance has been evaluated on the fully three-dimensional example of helicoidal spring cantilever subjected to the continuous constant force load [2] (see figure 5). three main aspects have been studied: numerical locking of timoshenko beam, comparison of convergence of the bernoulli and timoshenko beam formulations, and error of the bernoulli formulation caused by the neglect of the shear effects. 5.1. numerical locking firstly, the removal of the numerical locking of the timoshenko beam formulation has been verified. the presented timoshenko formulation without and with locking treatment have been compared using the example of the thin helix depicted in figure 5. the convergence of the tip vertical displacement has been observed. additionally, the results have been compared to the standard fem beam element with cubic approximation, which is naturally locking-free [5]. as expected, the numerical locking can be observed when no locking treatment is used. the convergence of the isogeometric beam is significantly deteriorated in comparison with the standard fem beam element. when b̄-locking treatment is used a superior convergence is obtained. the slow convergence of the traditional fem beam is due to the inability to exactly represent curved geometry. see figure 6 for the results. on the other hand, isogeometric formulation with the locking treatment provides superior convergency. 5.2. convergence comparison in order to compare the timoshenko and the bernoulli beam formulations the error of the vertical tip displacement has been calculated as e = wref − w wref . (13) the reference solution have been calculated using fine mesh consisting of 320 control points individually for each formulation. the results are shown in figure 7. the better convergence is provided by the presented bernoulli beam formulation, nevertheless, the timoshenko beam element performs also very well. 5.3. applicability of bernoulli formulation finally, the applicability of the bernoulli beam formulation has been assessed. the reference solution of the timoshenko beam formulation has been used r1 r2 helicoidal spring: r = 2 m h = 3 m load: f = 3 kn/m cross-section: r1 = 0.11 m r2 = 0.15 m material: e = 200 gpa figure 5. helicoidal spring cantilever subjected to the constant force load and its cross-section. in order to verify the error caused by the neglect of the shear effects. the analysis has been performed for five different cross-sections in order to assess the behavior of beams with various thickness/length ratios. the results can be observed in table 1. as expected, the error rises significantly with the increasing radius of the cross-section. r1 0.11 0.31 0.51 0.71 0.91 r2 0.15 0.35 0.55 0.75 0.95 r2/l 0.012 0.027 0.042 0.058 0.074 error 0.7% 4.5% 10.9% 18.7% 27.1% table 1. the error of bernoulli beam formulation to the timoshenko reference solution for different crosssections. 6. conclusions two isogeometric beam elements have been implemented: the bernoulli beam element based on the work of a. m. bauer et al. [1] and the timoshenko beam element presented by g. zhang et al. [2] with b̄method [6] used as a locking treatment. three main aspects have been observed in order to compare the performance of the presented formulations. firstly, the numerical locking of the timoshenko beam element has been documented and the suitability of the locking treatment used has been verified. secondly, the convergence properties of both 15 edita dvořáková, bořek patzák acta polytechnica ctu proceedings -0.023 -0.0229 -0.0228 -0.0227 -0.0226 -0.0225 -0.0224 -0.0223 20 40 60 80 100 120 140 160 ve rti ca l t ip d is pl ac em en t number of control points (nodes) ref timoshenko timoshenko bbar fem -0.02293 -0.02292 -0.02291 -0.02290 -0.02289 -0.02288 -0.02287 20 40 60 80 100 120 140 160 ve rti ca l t ip d is pl ac em en t number of control points (nodes) ref timoshenko bbar fem figure 6. comparison of timoshenko beam element with no locking treatment with timoshenko beam element using b̄-method and with naturally locking-free standard fem beam element. the vertical axis corresponds to the vertical tip displacement w, the horizontal axis corresponds to the number of nodes (control points). 16 vol. 30/2021 comparison of timoshenko and bernoulli beams 0 1x10-6 2x10-6 3x10-6 4x10-6 5x10-6 6x10-6 7x10-6 8x10-6 9x10-6 30 35 40 45 50 55 60 65 70 75 e rr or o f t he v er tic al ti p d is pl ac em en t number of control points (nodes) bernoulli timoshenko bbar figure 7. comparison of covergence of timoshenko beam element and bernoulli beam element. the vertical axis corresponds to the erro of vertical tip displacement w, the horizontal axis corresponds to the number of nodes (control points). the error is calculated with respect to the reference solution of particular beam theory. formulations have been analyzed, proving the better behavior of the bernoulli formulation. finally, the applicability of the bernoulli beam element only to the thin structures has been demonstrated. the superior convergence compared to the traditional fem formulation can be observed for both formulations. acknowledgements the financial support of this research by the grant agency of the czech technical university in prague (sgs project no. sgs20/038/ohk1/1t/11) is gratefully acknowledged. references [1] a. bauer, m. breitenberger, b. philipp, et al. nonlinear isogeometric spatial bernoulli beam. computer methods in applied mechanics and engineering 303:101–127, 2016. [2] g. zhang, r. alberdi, k. khandelwal. analysis of three-dimensional curved beams using isogeometric approach. engineering structures 117:560–574, 2016. [3] t. j. r. hughes, j. a. cottrell, y. bazilevs. isogeometric analysis: cad, finite elements, nurbs, exact geometry and mesh refinement. comput methods appl mech engrg 194:4135–4195, 2005. [4] l. piegl, w. tiller. the nurbs book. springer-verlag berlin heidelberg, new york, 1997. [5] z. bittnar, j. šejnoha. numerické metody mechaniky 1. české vysoké učení technické, praha, 1992. [6] r. bouclier, t. elguedj. locking free isogeometric formulations of curved thick beams. comput methods appl mech engrg 245-246:144–162, 2012. 17 acta polytechnica ctu proceedings doi:10.14311/app.2019.23.0054 acta polytechnica ctu proceedings 23:54–57, 2019 © czech technical university in prague, 2019 available online at https://ojs.cvut.cz/ojs/index.php/app a comparison of testing methods for determination of sprayed concrete tensile strength martin závacký brno university of technology, faculty of civil engineering, institute of geotechnics, veveří 331/95, 602 00 brno, czech republic correspondence: zavacky.m@fce.vutbr.cz abstract. sprayed concrete is important construction material in tunnelling. primary lining is essential in natm where the sprayed concrete can be loaded by tension due to bending moments. the tension is common reason of failure because concrete has a relatively low tensile strength. the tensile strength is usually determined by splitting tensile test in laboratory. however, the results can be distorted because the specimen is not loaded by pure tension in this case. the paper compares results of concrete tensile strength determined by two methods: indirect by the splitting tensile test and direct by the modified tensile test. keywords: tensile strength, sprayed concrete, modified tensile test, splitting tensile test. 1. introduction sprayed concrete is an important construction material, especially in underground supporting structures. thus, knowledge of the mechanical properties of the material is necessary for project design. unconfined compressive strength is a very common and easily managed testing method. however, tensile strength is also a significant parameter. the tensile strength of concrete is relatively low compared to the compressive one, and therefore failure of the material often occurs through the action of tensile forces. direct tensile strength testing comes with several complications such as the requirement of special testing equipment, a high precision in specimen preparation and especially the gripping of a specimen during testing. these identified problems are often overcome by using various indirect tensile testing methods, e.g. a splitting tensile test (stt) or a flexural test. however, indirect methods may result in different values of tensile strength and they often overestimate parameter values [1]. comprehensive comparative studies have been carried out also in the field of rock mechanics [2, 3]. thus, direct tensile testing is still important. the modified tension test (mtt) was suggested in order to eliminate the previously stated technical problems [4] and the method has already been well established in practice e.g. at graz university of technology [5]. tension is induced by special geometry of a specimen (figure 1), therefore a minimal adaptation of the commonly used compression apparatus is needed. the main benefit of the method is the entire elimination of specimen gripping problems. the main objective of this presentation is to compare the tensile strength of sprayed concrete obtained by two different methods. a series of splitting tensile tests and modified tensile tests were carried out with figure 1. modified tension test setup [8]. the results showing variation between the methods. the study also extended previous experience of the authors with laboratory testing of sprayed concrete [6], which can be employed for example in the determination of input parameter values for the shotcrete model [7]. 2. materials and methods the concrete mixture used for preparation of specimens is described in table 1. one block of sprayed concrete was used for extraction of all specimens by core drilling (figure 2). a set of 8 specimens with diameter of 55 mm and length to diameter ratio 1.0 were used for the splitting tensile test (stt) and a set of 3 specimens for the modified tension test (mtt). the mtt specimens were prepared with an outer diameter of 100 mm and a length to diameter ratio of 1.5. the inner drilling diameter was 43 mm and the middle one was 69 mm. the extraction and further 54 https://doi.org/10.14311/app.2019.23.0054 https://ojs.cvut.cz/ojs/index.php/app vol. 23/2019 determination of sprayed concrete tensile strength figure 2. scheme of specimens extraction. loading of the specimens gave attention to orientation of the failure plane. thus, the failure plane was in all cases parallel to the layering of sprayed concrete. the tests were carried out by servo-hydraulic compression load frame under load control and a rate of loading 200 n/s in both the used methods. the maximal force reached was recorded and tensile strength was calculated according to the following formula: σstt = (2fmax) / (πdl) (1) for the splitting tensile test, where fmax – maximal force; d – specimen diameter; l – specimen length σmtt = fmax/ ( πr21 −πr 2 2 ) (2) for the modified tension test, where there is fmax – maximal force; r1 – radius of larger core hole (middle drilling); r2 – radius of smaller core hole (inner drilling). the dry unit weight of specimens for stt was determined as well, in order to evaluate the material homogeneity. the obtained values of dry unit weight and σstt for each specimen were assessed by the kolmogorov-smirnov test (k-s test) if the data has a normal distribution. the one-sample k-s test is a nonparametric test of the equivalence of a continuous probability distribution that can be used to compare a sample with a reference probability distribution [9]: dα ≥ dn = max{ ∣∣f (x(i)) − (i− 1) /n∣∣ ,∣∣f (x(i)) − i/n∣∣}, i = 1, 2, ...,n (3) where there is dα – critical value (significance level α= 0.05 was used); f(x) – reference probability distribution (considered as normal distribution); n – number of samples. the splitting tensile tests were carried out with respect to testing standard requirements čsn en 123906. the modified tension test procedure was carried out according to suggestions and reported experience published in previous studies [3, 8]. the tension annulus in the specimens was prepared with a thickness of 13 mm. the maximal aggregate diameter was 8 mm, and hence the obtained tensile strength should not be influenced by oversized aggregates. 3. results results of the splitting tensile tests carried out according to the method described in the previous chapter are listed in table 2. the table contains also data of dry unit weight and statistical parameters as averages and standard deviations. results of modified tension tests are listed in table 3. the dry unit weight (table 2) showed low variation when the maximal value (2247 kg/m3) is only 1.3 % above the average and the minimal value (2175 kg/m3) is only 1.9 % below the average. a standard deviation of σstt (0.33 mpa → 8 %) is more significant than in the case of dry unit weight (table 2). however, both parameters appeared as relatively consistent, and thus the tested material can be considered as homogeneous. vice-versa, the mtt results showed up as non-consistent, due to a significant drop of tensile strength in the case of specimen 2 (table 3). eccentricity of drilling was found by investigation of the specimen after the test (figure 4). hence, distribution of load in the specimen was not uniform during the test and concentration of stress probably caused failure under the lower level of loading [3]. therefore, only the results of tests 1 and 3 were considered as valid. in comparison with the stt average, mtt tensile strength was determined as 28 % lower. plotting of the cumulative distribution function against data of dry unit weight and stt tensile strength are presented in figure 3. in both cases dn value of k-s test was lower than the critical value dα (see formula 3). therefore, the normal distribution of the data was judged as appropriate. 4. conclusions a series of stt and mtt tests were carried out with 28 % higher value of tensile strength by stt. thus, overestimation of tensile strength by the indirect method was indicated, as was stated in previous studies [3, 8]. the main technical problems of direct tensile testing are eliminated by mtt. however, successful application of the method remains strongly dependent on the precision of specimen preparation. this fact was investigated in [3]. and was also observed during the conducted tests (figure 4). uncertainty about stress distribution within the geometry of mtt specimens was noted by molenda [3], and thus additional research from a mechanical point of view would be beneficial. the experience obtained during the presented experiments may be useful in further laboratory testing of either shotcrete or rocks regarding the purpose of determining input parameter values for various material models used in geotechnical analysis. acknowledgements this paper has been researched and written under the project no. lo1408 "admas up – advanced materials, structures and technologies", supported by the ministry of 55 martin závacký acta polytechnica ctu proceedings cement water/cement ratio aggregate 0/1 aggregate 0/4 aggregate 4/8 500 kg/m3 0.4 174 kg/m3 919 kg/m3 585 kg/m3 table 1. concrete mixture specification. specimen % [kg/m3] σstt [mpa] statistic 1 2214 3.97 average % 2 2234 4.28 2217 3 2230 4.13 std % 4 2231 3.63 24 5 2247 3.83 average σstt 6 2218 4.38 4.15 7 2175 4.66 std σstt 8 2189 4.33 0.33 table 2. splitting tensile test and dry unit weight results. specimen σmtt [mpa] 1 3.13 2 2.25 3 3.28 average (1 and 3) 3.21 table 3. modified tension test results. figure 3. normal distribution test: left – dry unit weight; right – stt tensile strength. vertical axis represents probability [-]. 56 vol. 23/2019 determination of sprayed concrete tensile strength figure 4. spec. 2 after mtt, detail of eccentricity. education, youth and sports under the "national sustainability programme i" and under the project no. fast-s18-5356 – "stanovení vstupních parametrů materiálových modelů pro potřeby podzemního stavitelství s možností využití optimalizačních metod" supported by brno university of technology – faculty of civil engineering. references [1] s. kang, j. park, g. ryu, et al. comparison of tensile strengths with different test methods in ultra high strength steel-fiber reinforced concrete. key engineering materials 417-418:649–652, 2010. [2] d. li, l. wong. the brazilian disc test for rock mechanics applications: review and new insights. rock mechanics and rock engineering 46:269–287, 2013. [3] m. molenda, f. stöckhert, s. brenne, m. alber. comparison of hydraulic and conventional tensile strength tests. effective and sustainable hydraulic fracturing pp. 981–992, 2013. [4] m. luong. a new test for tensile strength measurements. revue francaise de géotechnique 34:69–74, 1986. [5] m. blümel. improved procedures for laboratory rock testing. in proceedings of conference eurock 2000. eurock, aachen, germany, 2000. [6] m. závacký, chalmovský, l. miča, p. bílek. stanovení hodnot vstupních parametrů pokročilého materiálového modelu pro stříkaný beton. in proceedings of conference zakládání staveb. česká geotechnická společnost čssi, 2017. [7] b. schädlich, h. schweiger. a new constitutive model for shotcrete. numerical methods in geotechnical engineering pp. 103–108, 2014. [8] r. plinninger, b. thomée. experimental and model studies on the modified tension test (mtt) – a new and simple testing method for direct tension tests. in proceedings of conference on geotechnical measurements and modelling. a.a. balmeka publishers, 2003. [9] kolmogorov-smirnov test. https://en.wikipedia. org/wiki/kolmogorov-smirnov_test. 57 https://en.wikipedia.org/wiki/kolmogorov-smirnov_test https://en.wikipedia.org/wiki/kolmogorov-smirnov_test acta polytechnica ctu proceedings 23:49–52, 2019 1 introduction 2 materials and methods 3 results 4 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.32.0015 acta polytechnica ctu proceedings 32:15–20, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague calculation of seismically loaded earth retaining structures jan pruška∗, veronika pavelcová czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague, czech republic ∗ corresponding author: jan.pruska@cvut.cz abstract. eurocode 1998-5 for the design of earth retaining structures in seismic conditions describes calculations using the mononobe-okabe method. this simplified quasi-static procedure is to some extent very conservative. the article describes the limits of this method and brings out the comparison using theoretical numerical examples with other approaches. the first, is a calculation method introduced in the american nhcrp 661, the second, is the chen and liu method and the last one is the mylokanis et al. method. keywords: earth retaining structures, seismicity, seismic coefficient, earth pressure, mononobe-okabe. 1. introduction the theory of coulomb and rankin has formed the basis for current commonly used theories of earth pressure calculation. during the 20th century, it was found that earth pressures during seismic action can be significantly higher than in the case of purely static loading conditions. okabe (1924) and mononobe & matsuo (1929) introduced the first method that could calculate the total earth pressure from static and seismic loading. this method became known as the mononobe okabe method (hereinafter m-o) and is recommended in many geotechnical standards. the eurocode 1998-5 for the design of retaining and gravity walls in seismic conditions also provides a calculation using this method. the m-o method as a simplified quasi-static procedure that is to some extent very conservative. the article describes the limits of this method, and finds comparison with other approaches for the calculation of seismically loaded retaining walls in the american nhcrp 661, also in the method of chen and liu (1990) and in an alternative to the standard m-o method given by mylonakis et al. 2. seismic loading for quasi-static calculations of gravity and retaining walls, the seismic soil load is given by horizontal and vertical forces which are equal to the weight of the active wedge w and the corresponding seismic coefficient kh, kv, see figure 1. the vertical load must be considered to have the most adverse effect. the value of the seismic coefficient is generally determined to be equal to the ratio of the considered acceleration and gravitational acceleration and it is possible to determine the relationships according to the different authors. figure 1. wall load during quasi-static calculation 2.1. according to terzaghi (1950) kh = ah g (1) kv = av g (2) where: ah design acceleration in the horizontal direction [m · s−2] av design acceleration in the vertical direction [m·s−2] g gravitational acceleration [m · s−2] 2.2. according to gutenberg and richter (1956) kh = a0 g (3) log10 a0 = −2.1 + 0.81ms − 0.027m 2 s (4) where: a0 earthquake acceleration with magnitude ms [m · s−2] 15 https://doi.org/10.14311/app.2022.32.0015 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en jan pruška, veronika pavelcová acta polytechnica ctu proceedings 2.3. according to eurocode 1998-5 kh = s r (5) kh = ±0.50kh if both av and g > 0.6 (6) kh = ±0.33kh for other cases (7) where: s the subsoil coefficient defined in eurocode 1998-1 [−] r coefficient according to table 1 [−] for walls lower than 10 m, the seismic coefficient must be constant in height. eurocode 1998-5 does not specify a procedure for determining the effect of higher walls on the value of seismic loading. 2.4. according to nchrp 611 and aashto the national cooperative highway research program (nchrp) is a research program of the united states of america. road infrastructure research is funded by the american association of state highway and transportation officials (aashto). report 611 presents current results from the research of seismic design of road infrastructure, including slopes and load-bearing structures. determination of seismic acceleration is not found in the nchrp 611 report itself but in superior document: lrfg bridge design specifications (aashto, 2007 and 2012). report 611 nchrp concerns this report in chapters 5 and 7. it introduces the factor α which takes into account the height of backfill. kh,nchrp = fpga pgaα g (8) where: kh,nchrp horizontal seismic coefficient kh,d as defined in aashto (2007/2012) [−] fpga locality coefficient for the zero period in the elastic response spectrum values in table 3.10.3.2-1 aashto (2012) [−] pga peak soil level acceleration for bedrock (class b) [m · s−2] α backfill height coefficient [−] g gravitational acceleration [m · s−2] the backfill height coefficient α described in the 611 nchrp report is the result of a one-dimensional free field wave propagation analysis. the analysis took into account the fact that the acceleration of the backfill of the supporting structure is not the same at any point. this coefficient can be calculated for soils of classes c, d and e (equivalent to all soils except rock and very soft clay) according to aashto as follows: α = 1 + 0.003 h [( 0.5 fv s1 fpga pga ) − 1 ] (9) h height of load-bearing structure or backfill [m] fv locality coefficient for long-term range in the elastic response spectrum according to table 3.10.3.2-3 aashto (2012) [−] s1 horizontal spectral acceleration coefficient for a period of 1 s in rock (class b) [m · s−2] however, equation (9) can only be used if the height of the supporting structure is h > 6 m. for the construction of a height h ≥ 30 m, the height in equation (9) must be set to h = 30 m. the backfill height coefficient corresponds to the function correction coefficient qh according to swiss sia 267. 3. mononobe-okabe method following the 1923 earthquake in kanto, japan, okabe (1924) and mononobe & matsuo (1929) performed a series of experiments on a vibrating table. a pseudostatic method (introducing the principle of equivalent accelerations) based on the obtained experimental results was derived to calculate the total earth pressure from both static and seismic loading. this method (known as the mononobe-okabe method), based on the coulomb’s limit approach and the assumption that the earth pressure is derived from a rigid body that moves along a shear plane, and gives relations for active or passive earth pressure under seismic loading as follows (the sign convention is in figure 2): pae = 1 2 γ (1 ± kv) h2kae (10) ppe = 1 2 γ (1 ± kv) h2kpe (11) kae = cos2(ϕ − θ − α) cos θ cos2 α cos(δ + θ + α) [ 1 + √ sin(ϕ + δ) sin(ϕ − β − θ) cos(β − α) cos(α + θ + δ) ]2 (12) kpe = cos2(ϕ − θ + α) cos θ cos2 α cos(θ − α) [ 1 − √ sin(ϕ) sin(ϕ + β − θ) cos(θ − α) cos(β − α) ]2 (13) 16 vol. 32/2022 calculation of seismically loaded earth retaining structures type of gravity or retaining wall r free gravitational walls that can deviate up to dr = 300α · s [mm] 2.0 free gravitational walls that can deviate up to dr = 200α · s [mm] 1.5 flexible concrete walls, anchored or spread walls, reinforced concrete walls 1.0founded on vertical piles, embedded foundation walls and bridge abutments table 1. values of the coefficient r for the calculation of the horizontal seismic coefficient (α is the ratio of the design acceleration ad for the foundation soil type a to the gravitational acceleration g, s is the subsoil coefficient defined in eurocode 1998-1) the following notation is used as in the previous equations: γ volumetric density of soil [kn · m−3] h height of load-bearing structure [m] kae active earth pressure coefficient (static + dynamic) [−] kpe passive earth pressure coefficient (static + dynamic) [−] α inclination of the back of the wall [◦] β backfill inclination behind the retaining wall [◦] δ angle of friction between the soil and the wall structure [◦] ϕ angle of internal friction of the backfill material [◦] kh seismic coefficient of horizontal acceleration [−] kv seismic coefficient of vertical acceleration [−] θ tan−1 ( kh 1 − kv ) [−] figure 2. sign convention (left: passive, right: active) in many cases, the seismic coefficient of vertical acceleration kv in these examples is neglected because the effect of vertical acceleration is relatively small. 3.1. limits of the mononobe-okabe method the relations of the m-o method are based on the following assumptions of the mohr-coulomb theory of soil failure with a criterion according to the coulomb theory: • dry, homogeneous, isotropic backfill material • the soil above the shear plane is solid it creates a wedge • the soil under the supporting structure is loadbearing • seismic load is modelled as a static force (multiple of kh,d) • the wall must be able to move or rotate so that a shear plane can form in the soil • the resulting shear plane is planar • the terrain above the wall is modelled as an infinitely long slope the last two items in this list are responsible for problems using the m-o method if the slope (backfill) behind the wall β reaches or exceeds the inclination of the shear plane, the result is an infinitely large "wedge" of soil. as a result, the calculated active earth pressure is infinitely large and the relations of the m-o method no longer provide a solution for these inclinations. proof of the limited validity of equation (12) can simply be shown by the example of determining the stability of an infinitely long slope with a planar shear plane where the slope is formed by incoherent soil and there is no water in it. the seismic load is given by the horizontal acceleration kh,d = ah,d/g and the vertical acceleration kv,d is neglected. it is known that such a slope is stable if the inclination β is less than the angle (ϕ−θ) or in the extreme case β = ϕ−θ when the shear plane is exactly parallel to the inclination of the surface. when using equation (12) for the active earth pressure coefficient, for kh,d = 0.1; β = 30◦, ϕ = 35◦, the expression sin(ϕ−β −θ) below the square root becomes negative and no natural solution can be found. another limitation is, that the m-o method has only limited suitability for calculating passive earth pressure from seismic loading because in many cases the shear plane at passive earth pressure does not have to be considered horizontal but logarithmic. if we compare the value of passive earth pressure using the logarithmic shape of the shear plane and the horizontal one (which the m-o method considers), it can be seen that a planar plane can cause significant overestimation of the resistance to passive earth pressure, especially when considering the angle between the soil and the wall structure δ = 2/3ϕ. the last disadvantage of the m-o method is the assumption of a constant seismic load over the entire height of the structure (i.e. the soil wedge). 17 jan pruška, veronika pavelcová acta polytechnica ctu proceedings 4. eurocode 1998-5 appendix e gives two formulas for calculating the active earth pressure coefficient kae. for the inclination of the terrain surface behind the wall β ≤ ϕ − θ uses the equation (12) according to m-o. but if the angle of the inclination behind the wall satisfies the condition β > ϕ−θ, the original relationship is modified by complete exclusion of the influence of the inclination of the terrain surface behind the wall β as follows: kae = cos2(ϕ − θ − α) cos θ cos2 α cos(δ + θ + α) (14) thus, for the inclinations of the terrain behind the wall with values of the angle β > ϕ − θ, a limit arises where the original relation (equation (12)) calculates the infinite coefficient of earth pressure kae but equation (14) gives a constant value of the coefficient kae. the inclination of the terrain behind the retaining wall can therefore be arbitrarily increased above the value ϕ − θ according to the equation (14) without increasing the soil pressure (see figure 3). in this way, the fact that the slope above the wall is already unstable and theoretically the whole area must be supported by the supporting structure so that its displacement does not occur can be missed. therefore, the use of the equation (14) is unsuitable in the case of very steep slopes of the terrain above the wall. the calculations for this paper have been carried out in the geo5 program which has, inter alia, a procedure according to eurocode 1998-5 built-in. the advantage of this program is that at the values of the inclination β > ϕ − θ it notifies the user to this fact and possible problems related to it. eurocode 1998-5 uses the equation (13) to calculate the passive earth pressure coefficient kpe. however, it is recommended to use this equation mainly for the vertical front side of the wall. in addition, it requires consideration of the friction between the wall and the soil δp = 0◦ (section 7.3.2.3 (6)) which leads to very conservative results. 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.1 1.2 0 5 10 15 20 25 30 35 40 45 k a e [] β [°] φ = 25; kh,d = 0,1 φ = 25; kh,d = 0,2 φ = 25; kh,d = 0,3 φ = 30; kh,d = 0,1 φ = 30; kh,d = 0,2 φ = 30; kh,d = 0,3 φ = 35; kh,d = 0,1 φ = 35; kh,d = 0,2 φ = 35; kh,d = 0,3 figure 3. dependence of the inclination of the terrain behind the wall and the coefficient kae according to eurocode 1998-5 5. chen and liu chen and liu’s method is based on chang’s doctoral dissertation (1981). as with the m-o method, it is based on a kinematic approach. however, it considers a logarithmic shear plane instead of a horizontal shear plane for the m-o method. chen and liu (1990) also present tabular values of earth pressure coefficients for the most common cases of retaning walls and backfill soils. this method generally provides lower passive earth pressures than the m-o method (for the same value of friction between wall and soil). 6. nchrp 611 the shamsabadi method (2007), recommended by nchrp 611, is a kinematic method that considers the logarithmic shear plane and the hyperbolic relationship between stress and strain. thus, this method differs from conventional kinematic methods such as chen and liu (1990). standard nchrp 611 (2008) in appendix 5 gives graphs for the passive earth pressure coefficient kpe as a function of variables: soil cohesion c, wall height h , soil volumetric density γ, seismic coefficient kh and soil internal friction angle ϕ. 7. mylonakis et al. in 2007, mylonakis et al. presented a closedform stress plasticity solution for gravitational and earthquake-induced earth pressures on retaining walls. the proposed solution is essentially an approximate yield-line approach, based on the theory of discontinuous stress fields. it is an alternative to the standard m-o equations for seismic earth pressures suffering thus from the same limits as the m-o method. on the other hand this alternative approach works with respect to active and passive conditions, as it can be expressed by a single equation with opposite signs for soil friction angle and wall roughness. according to this method the active and passive presure under seismic loading can be determined by: pe = 1 2 γ (1 − kv) h2ke (15) where ke = cos(α − β) cos(β + θ) cos θ cos δ cos2 α · · [ 1 ∓ sin ϕ cos(∆2 ∓ δ) 1 ± sin ϕ cos [(∆1 ± (β + θ)] ](∓2ψe tan ϕ) (16) the caquot angles ∆1, ∆2 are given by mathematical expressions sin ∆1 = sin β sin ϕ (17) sin ∆2 = sin δ sin ϕ (18) 18 vol. 32/2022 calculation of seismically loaded earth retaining structures and twice the revolution angle of principal stresses in the two regions under seismic conditions 2ψe = ∆2 ∓ (∆1 + δ) + β − 2α − θ (19) the upper signs in the equations refers to active conditions pae and the lower signs to passive ppe and the notation used in equations above is the same as in the standard m-o method (section 3). 8. comparison of methods with the active seismic earth pressure coefficient to compare the methods of chen and liu, mylonakis at el. and m-o with the modification given by the equation (14) for backfill inclinations β > ϕ − θ according to the eurocode 1998-5, a simple case was deliberately chosen: a gravity wall with a vertical back side of the wall (α = 0◦) with a height h = 3 m and a wall friction angle δ = 2/3ϕ, inclination of the terrain surface β = 0–30◦, soil friction angle ϕ = 30◦, cohesion c = 30 kpa. the vertical acceleration of the earthquake was neglected, the horizontal acceleration kh,d is 0.1 and 0.2, respectively (see figure 4). � � �a h p ae w khdw figure 4. scheme of solved example the differences between methods are minor for higher inclinations of the terrain surface β (figures 5). the reason for this is the fact that the logarithmic shear plane considered by chen and liu’s method becomes practically flat for the calculations of the active earth pressure coefficient performed here. as mentioned before, it can be observed that mylonakis’s et al. method is not able to give any natural solution after exceeding the critical backfill inclination β. the peculiarity of the chen and liu’s (1990) method is evident from both graphs in figure 5 where it can be seen that the method gives results even for backfill inclinations above the critical inclination β although for very steep slopes these values are already above the physically permissible limit (and the slope would be unstable). signs of chen and liu’s logarithmic shear plane are visible only in figure 5 (bottom) as the values of the active earth pressure coefficient slightly differs from other methods. 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.1 1.2 0 5 10 15 20 25 30 k a e [] β [°] mononobe-okabe chen & liu mylonakis et al. 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.1 1.2 0 5 10 15 20 25 30 k a e [] β [°] mononobe-okabe chen & liu mylonakis et al. figure 5. comparative calculations of the effect of cohesion on active seismic earth pressure coefficient: kh = 0.1 (top) and kh = 0.2 (bottom). 9. comparison of methods with the passive seismic earth pressure coefficient to compare the values of the passive earth pressure coefficient kpe according to chen and liu (1990) and eurocode 1998-5 (figure 6) a gravity wall with a vertical back side of the wall (α = 0◦), with a height h = 3 m and horizontal backfill surface behind the wall β = 0◦ was chosen. not only the standard m-o method, as the eurocode 1998-5 recommends, but also the alternative approach of the m-o method presented by mylonakis et al., (as it should theoretically give exactly the same results as the standard m-o method), was used for the calculation. for the chen and liu (1990) method, the wall friction angle equals δ = 2/3ϕ, for eurocode 1998-5 (m-o and mylonakis et al.) δ = 0◦ (prescribed by the standard), the internal friction angle ϕ was considered from 25◦ to 35◦. the vertical acceleration of the earthquake was neglected, the horizontal acceleration of kh,d took values of 0, 0.1, 0.2 and 0.3. 19 jan pruška, veronika pavelcová acta polytechnica ctu proceedings 0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 8.0 9.0 25 27 29 31 33 35 k p e [] ϕ [°] m-o khd = 0.0 m-o khd = 0.1 m-o khd = 0.2 m-o khd = 0.3 ch&l khd = 0.0 ch&l khd = 0.1 ch&l khd = 0.2 ch&l khd = 0.3 m. et al. khd = 0.0 m. et al. khd = 0.1 m. et al. khd = 0.2 m. et al. khd = 0.3 mononobe-okabe δp = 0° chen & liuδp = -2/3 ϕ mylonakis et al.δp = 0° figure 6. comparison of methods for calculating the passive coefficient of seismic earth pressure according to the literature, the procedures recommended by eurocode 1998-5 give (δ = 0◦) (which corresponds to its requirements) the most conservative values of all the methods. in fact the method of mylonakis et al. gives even lower values of passive pressure coefficient than the standard m-o method but taking into account the scale of the values of chen and liu’s method, the difference is insignificant. if the value δ = −2/3ϕ was used for the friction between the wall and the soil using the standard m-o method, higher coefficients of passive earth pressure would be obtained than by following the chen and liu’s method (1990). however, since many derivations of the passive seismic pressure coefficient kpe reported in the literature are controversial, it is recommended to use the calculation involved in the m-o method according to eurocode 8 (δp = 0◦). 10. conclusion it should be noted that the m-o method represents a considerable simplification of reality. in addition to the limitations of all pseudo-static methods, it should be emphasized that the frequency range of excitation or soil stiffness are not taken into account in seismic loading while using the m-o method. in the case of solving the problem with the value of the inclination of the terrain β, we obtain constant values of the coefficient of active earth pressure kae for and above critical inclination angle β using eurocode 19985 and while determinating the passive earth pressure according to eurocode 1998-5, the requirement of friction between the wall and the soil (δp = 0◦) (section 7.3.2.3) has to be satisfied. in both of the mentioned cases, it seems more advantageous to use the chen and liu’s method. another disadvantage is that when calculating walls higher than 10 m, eurocode 1998-5 does not specify the procedure for determining the seismic load (unlike the american nchrp 611). it also cannot be overlooked that the classical m-o equations are empirical and hardly derived by physicbased reasoning. on the other hand, the alternative approach presented by mylonakis et al. proposes a simpler solution in which the main features of the mathematical expressions, including signs, can be deduced by such reasoning from physics. moreover, the proposed solution is symmetric with respect to active and passive conditions, as it can be expressed by a single equation with opposite signs for soil internal friction angle ϕ and angle of friction between the soil and the wall structure δ. despite all the limitations and possible shortcomings, the m-o method is still used as an easy-tounderstand and meaningful method for the calculation of seismic loaded earth retaining structures. also eurocode 1998-5: eurocode 8: design of structures for earthquake resistance part 5: foundations, retaining structures and geotechnical aspects explicitly refers to this method in appendix e. acknowledgements the support provided by the sgs project no. sgs21/045/ohk1/1t/11. references [1] aashto. lrfd bridge design specifications, vol. 4th edition. american association of state highway and transportation officials, 2007. [2] eurocode 8. : design of structures for earthquake resistance–part 5: foundations, retaining structures and geotechnical aspects en 1998-5: 2004. brussels: european committee for standardization 2006. [3] fhwa (federal highway administration). seismic retrofitting manual for highway structures. part 1, bridges. tech. rep., turner-fairbank highway research center, 2006. [4] geo 5. user manual. http://www.fine.cz. [5] b. gutenberg, c. richter. magnitude and energy of earthquakes. annali di geofisica 9:1–15, 1956. [6] w.-f. chen, x. liu. limit analysis in soil mechanics. elsevier, 1990. [7] n. mononobe. on determination of earth pressure during earthquake. in proc. world engineering congress, vol. 9, pp. 177–185. 1929. [8] nchrp 611. seismic analysis and design of retaining walls, buried structures, slopes, and embankments. 2008. [9] s. okabe. general theory of earth pressure. journal of thejapanese society of civil engineers-1926-12 (1) 311, 1926. [10] pruška. dynamika geotechnických konstrukcí. čvut fsv, 2017. [11] a. shamsabadi. three-dimensional nonlinear seismic soil-abutment-foundation-structure interaction analysis of skewed bridges. university of southern california, 2007. [12] k. terzaghi. mechanism of landslides. application of geology to engineering practice, geol soc am pp. 83–123, 1950. [13] g. mylonakis, p. kloukinas, c. papantonopoulos. an alternative to the mononobe–okabe equations for seismic earth pressures. soil dynamics and earthquake engineering 27(10):957–969, 2007. 20 http://www.fine.cz acta polytechnica ctu proceedings 32:15–20, 2022 1 introduction 2 seismic loading 2.1 according to terzaghi (1950) 2.2 according to gutenberg and richter (1956) 2.3 according to eurocode 1998-5 2.4 according to nchrp 611 and aashto 3 mononobe-okabe method 3.1 limits of the mononobe-okabe method 4 eurocode 1998-5 5 chen and liu 6 nchrp 611 7 mylonakis et al. 8 comparison of methods with the active seismic earth pressure coefficient 9 comparison of methods with the passive seismic earth pressure coefficient 10 conclusion acknowledgements references doi:10.14311/app.2021.30.0098 acta polytechnica ctu proceedings 30:98–103, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app optimization of tuned mass dampers attached to damped structures minimization of maximum displacement and acceleration jan štěpánek∗, jiří máca czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jan.stepanek@fsv.cvut.cz abstract. a tuned mass damper is a device, which can be highly helpful while dealing with dynamic behaviour of structures. its proper design is conditioned by knowledge of both loading and the structure properties. in many cases, the structure can be represented by single degree of freedom model, which simplifies the design and optimization of tuned mass dampers. most of studies focus only on minimization of displacement of the main structure under harmonic force load, however, in many cases, different frequency response function would be more appropriate. this paper presents an extension of design formulas for the h∞ optimization of tuned mass dampers for damped structures and various frequency response functions. keywords: acceleration, damped structure, displacement, dynamic response, optimization, tuned mass damper. 1. introduction optimal parameters of a tuned mass damper (tmd) have been a subject of many researches since the first half of the 20th century. since then, several different optimization criteria have been proposed. the most common ones are called h2, which minimizes the area under a frequency response function (frf) and h∞, which minimizes the maximum amplitude magnification factor. a first set of nearly optimal tmd parameters for undamped structure came from the existence of invariant points, where the amplitude is independent of tmd damping [1, 2]. a closed form solution of h∞ optimum parameters of tmd attached to a damped structure have not been found. in 1978, ioi and ikeda presented empiric design formulas [3], however, they were derived only for an insufficiently large mass ratio µ > 0.03, which makes them inappropriate for certain civil engineering structures. it is interesting that despite the formulas being derived for a large mass ratio, they are recommended in footbridge design guidebook setra [4]. design charts allowing use of lower mass ratio µ > 0.01 were published in 1981 [5]. asami et al. used a perturbation method to modify the fixed-point solution regarding a main system damping ratio as a perturbation [6]. they obtained a series solution, which is, however, too complicated for engineering application. therefore, they have proposed a simplified version. abubakar presented a simple modification of the fixed-point tmd optimum for damped structure. this modification provides very accurate solution for minimization of displacement [7]. despite many useful formulas and design methods being proposed, some modern guidebooks for civil engineering, for example [4, 8], keeps using the original approximate solutions thanks to their easy application. most of studies focus on a reduction in a maximum amplitude of displacement. this is a suitable criterion for the reduction of structural stress due to a dynamic load. however, in many cases, the main purpose of tmd is to reduce the maximum amplitude of acceleration of the main structure. these cases contain machines and sensors which are sensitive to vibration, because the vibration tolerance is usually defined in the units of acceleration. furthermore, human comfort criteria inside buildings or on footbridges are usually limited by acceleration. this paper points out the importance of distinction between displacement and acceleration criteria of the h∞ optimization. the approximate optimal stiffness and damping ratio for both criteria are proposed in this paper covering both ground and force excitation. considering the importance of easy practical application of the result, the formulas for optimal parameters are composed of standard solutions given by the fixed-point method and a correction, which takes nonzero damping into account. therefore, in the case of zero structural damping, we obtain standard result given by the fixed-point method. in the case of nonzero damping, we obtain results given by the modified formulas. 2. equations of motion the main structure is represented by a single degree of freedom (sdof) system and it consists of a mass m1, stiffness k1, and damping c1. tmd is composed of a mass m2, an elastic member with stiffness k2 and a viscous member with damping c2. the struc98 http://dx.doi.org/10.14311/app.2021.30.0098 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 optimization of tuned mass dampers in damped structures figure 1. a) force excitation, b) support excitation. ! k1 + k2 −k2 −k2 k2 " # x1 x2 $ + ! c1 + c2 −c2 −c2 c2 " # ẋ1 ẋ2 $ + ! m1 0 0 m2 " # ẍ1 ẍ2 $ = f (t) (1) ture can be loaded by a harmonic force (figure 1a) or excited by support motion (figure 1b). this paper covers both loading cases, because the optimal parameters of tmd are not identical for force and support excitation. the equations of motion are described by eq. 1, where f (t) = # f (t) 0 $ in the case of force excitation, and f (t) = # −ẍ0(t)m1 −ẍ0(t)m2 $ in the case of support excitation. x, ẋ and ẍ denote the displacement, velocity and acceleration, respectively. it is important to note that x induced by support excitation is a relative displacement between the support and the mass. the absolute displacement is xa = x + x0. 3. optimal parameters in order to solve equation 1, it is advantageous to use following dimensionless substitutions and notation for further operations [9]: ω1 = % k1/m1 natural frequency of the main structure, ω2 = % k2/m2 natural frequency of tmd, µ = m2/m1 mass ratio, β = ω2/ω1 frequency ratio, ω = ω/ω1 forcing frequency ratio, c1,cr = 2 √ k1m1 critical damping of the main structure, c2,cr = 2 √ k2m2 critical damping of tmd, ξ1 = c1/c1,cr damping ratio of the main structure, ξ2 = c2/c2,cr damping ratio of tmd, nf (ω)= (2ξ2βω)2 + (β2 − ω2)2, na(ω)= [1 + (2ξ1ω)2][(β2 − ω2)2 + (2ξ2βω)2], nr (ω) = [(1 + µ)β2 − ω2] 2ω4 + (1 + µ)2(2ξ2βω)2ω4, d(ω) = [(1 − ω2)(β2 − ω2) − µβ2ω2 − 4ξ1ξ2βω2] 2 + +4ω2[(β2 − ω2)ξ1 + (1 − ω2 − µω2)βξ2] 2. assuming a harmonic load according to figure 1, a steady state part of response can be expressed from eq. 1. the frequency response functions which describe the normalized response of the structure are summarized in table 1. 3.1. fixed-point approach the approximate h∞ optimum for main structure with no or negligible damping can be reached by the fixed-point method. this method was firstly described by den hartog [1]. he postulated that the approximate optimum frequency ratio βuopt is reached if the value of frf is equal in invariant points p and q, where the value of frf is not affected by tmd damping c2. brock lately stated that the optimum damping ratio ξu2,opt can be taken as the average of two optima, each calculated separately for one of two points p and q and he derived the well known formula of optimum damping for frf no. 1 in table 1 [2]. using this approach, we can derive the approximate optimum solutions for various frfs. the results are summarized in table1. it should be noted that the optimum parameters are identical for transfer functions 1 and 3, as well as for 2 and 4. this is can be seen in na(ω, ξ1 = 0) = nf (ω). however, this is true only if ξ1 = 0. 3.2. non-zero damping the fixed-point approach is not available if we introduce damping of the main structure. an analytical solution of optimum tmd parameters attached to a damped main structure have not been found and according to asami et al., the solution is probably impossible [6]. therefore, we need to rely on empiric and approximate solutions. in order to define an empiric formulas for optimum frequency ratio βopt and optimum damping ratio ξopt,2, the values of the parameters were found numerically on a surface of 44x41 points for 0.001 ≤ µ ≤ 0.2 and 0 ≤ ξ1 ≤ 0.2. the density of mesh was increased in the area of low mass ratio 0.001 ≤ µ ≤ 0.005, because in this range, the optimum values of βopt and 99 jan štěpánek, jiří máca acta polytechnica ctu proceedings no. excitation frequency responsefunction type frequency response function β u opt ξ u 2,opt 1 force displacement &&&& x1 f0/k1 &&&& = ' nf (ω) d(ω) 1 1 + µ ' 3µ 8(1 + µ) 2 force acceleration &&&& ẍ1 ω21 f0/k1 &&&& = ω 2 ' nf (ω) d(ω) ( 1 1 + µ ' 3µ 4(2 + µ) 3 support abs. displacement &&&& xa1 x0 &&&& = ' na(ω) d(ω) 1 1 + µ ' 3µ 8(1 + µ) 4 support abs. acceleration &&&& ẍ1 a ω21 x0 &&&& = ω 2 ' na(ω) d(ω) ( 1 1 + µ ' 3µ 4(2 + µ) 5 support rel. displacement &&&& ω21 x1 ẍ0 &&&& = 1 ω2 ' nr (ω) d(ω) ' 2 − µ 2(1 + µ)2 ' 3µ 4(1 + µ)(2 − µ) 6 support rel. acceleration &&&& x1 x0 &&&& = ' nr (ω) d(ω) ' 2 + µ 2(1 + µ)2 ' 3µ 8(1 + µ) table 1. optimum parameters of tmd based on the fixed-point method. ξopt,2 change more rapidly then in the rest of the function. a convergence problem may occur while finding ξopt,2 numerically, because a structural response is much more sensitive to tmd tuning than to its damping. the empiric formulas are combination of the fixedpoint solution and a correction function, which takes the structural damping into account. the original approach [3] uses a correction function composed of 2nd order polynomial of µ and ξ1, but we found it neither simplest nor best-fitting. numerous types of functions, including polynomial, exponential, logarithmic and hyperbolic ones, were tested using a non-linear regression to find the best approximation of the optimum tmd setting regarding a simple application. thanks to similar characteristics of the optimum parameters for all frequency response functions, two common function were defined: βopt = βuopt + ξ1 ) a1ξ1 + a2 * µ 1 + µ + 1 3 , (2) ξ2,opt = ξu2,opt + ξ1[b1 + b2 ln(µ) + b3µ + b4µξ 2 1 ] (3) as we can see in the shape of eqs. (2) and (3), the optimum values of tmd parameters βopt = βuopt and ξ2,opt = ξu2,opt if ξ1 = 0. in other words, the formulas preserve the fixed-point approach if the main structure remains undamped. the coefficient of determination was chosen as the comparative criterion for function fitting and it reached higher than 0.997 in all cases of frfs for our selected function types. the maximum error of βopt was lower than 1%. the maximum error of ξ2,opt reached up 5%, but it occurred only for extremely small mass ratio, where the optimum damping is close to zero and such a high error may be produced rather by a precisions of the numerical solution than by an inaccurate approximation. the average relative error of ξ2,opt was lower than 1% for all transfer functions. the coefficients ai and bi were found using linear regression and they are summarized in table 2. figure 2 presents the improvement of frf using desing formulas (2) and (3) for relative displacement of structure under support excitation. figure 2. improvement of tf 5 (µ = 0.05, ξ1 = 0.05) using desing formulas (2) and (3). figures 3-8 compare the numerical results with the approximations given by design formulas. we can see that in most cases, the empiric formulas give very precise results. it is also visible that ξ2,opt rises as the structural damping increases. however, the effect of ξ1 to the optimum frequency ratio is slightly more 100 vol. 30/2021 optimization of tuned mass dampers in damped structures frf no. a1 a2 b1 b2 b3 b4 1 -0.7636 -0.8748 0.1801 0.0140 0 -4.6350 2 1.2470 0.2644 0.2101 0.0178 1.264 0 3 -0.5561 -0.8466 0.1782 0.01367 0.01865 -4.3310 4 1.5600 0.2778 0.1985 0.01552 1.1420 29.0900 5 -0.7639 -1.5300 0.2005 0.01650 0.4629 0 6 0.9731 -0.2176 0.1783 0.0145 0 0 table 2. coefficients ai and bi for design equations (2) and (3). figure 3. optimum tmd parameters: tf no. 1. figure 4. optimum tmd parameters: tf no. 2. figure 5. optimum tmd parameters: tf no. 3. 101 jan štěpánek, jiří máca acta polytechnica ctu proceedings figure 6. optimum tmd parameters: tf no. 4. figure 7. optimum tmd parameters: tf no. 5. figure 8. optimum tmd parameters: tf no. 6. complicated. as we can see in figures 3-8, increase of ξ1 leads to decrase of βopt in the case of displacement frfs (no. 1,3,5). in the case acceleration frfs (no. 2,4,6), the increase of ξ1 has an opposite effect. sensitivity of optimum parameters to the damping ratio ξ1 varies among frfs. it is interesting to note that in the case of acceleration frfs, the optimum tuning βopt may slightly rise above one, which is an unusual situation in tmd design. 4. conclusions the design formulas for h∞ optimization of tmd attached to a damped structure are presented in this paper. the formulas are appropriate for tmds with the mass ratio µ between 0.001 and 0.2, and for the damping ratio of the main structure between 0 and 102 vol. 30/2021 optimization of tuned mass dampers in damped structures 0.2. this is a sufficient range for most of tmds localized in civil engineering structures and we consider the error of design parameters negligible in this range. the formulas were optimized for six frequency response functions, which describe relative and absolute displacement or acceleration for both force and support excitation. it is necessary to chose a suitable frf for tmd optimization regarding the purpose of tmd. for the purpose of a stress reduction of the main structure, frfs which describe displacement should be used. on the other hand, in order to increase a comfort of people in high-rise buildings and pedestrian on footbridges, or to reduce the vibration of equipment localized in the main structure, the acceleration frfs are an appropriate choice. we also find out that reduction of acceleration may lead to tmd tuned to a resonant or even slightly higher frequency. an easy application is preserved. only two simple formulas are presented, however, they can be used for all considered transfer functions with different parameters ai and bi. acknowledgements the authors gratefully acknowledge support from the czech technical university in prague, project sgs20/038/ohk1/1t/11 development and application of numerical algorithms for analysis and modelling in mechanics of structures and materials. references [1] j. p. den hartog. mechanical vibrations. mcgraw hill, new york, 1934. [2] j. brock. a note on the damped vibration absorber. journal of applied mechanics 68(a):284, 1946. [3] t. ioi, k. ikeda. on the dynamic vibration damped absorber of the vibration systems. bulletion of the jsme 21(151):64–71, 1978. [4] setra. footbridges: assessment of vibrational behaviour of footbridges under pedestrian loading, 2006. [5] s. e. randall, d. m. halsted, d. l. taylor. optimum vibration absorbers for linear damped systems. journal of mechanical design 103:908–913, 1981. [6] t. asami, o. nishihara, a. m. baz. analytical solutions to h∞ and h2 optimization of dynamic vibration absorbers attached to damped linear systems. journal of vibration and acoustics 124:284–295, 2002. [7] i. m. abubakar, b. j. m. farid. generalized den hartog tuned mass damper system for control of vibrations in structures. earthquake resistant engineering structures 104:185–193, 2009. [8] fib bulletin 32: guidelines for the design of footbridges, 2005. [9] o. nishihara, t. asami. closed-form exact solution to h infinity optimization of dynamic vibration absorber: ii. development of an algebraic approach and its application to a standard problem. proceedings of the spie, 3989, 2000. doi:10.1117/12.384589. 103 https://doi.org/10.1117/12.384589 acta polytechnica ctu proceedings doi:10.14311/app.2019.24.0044 acta polytechnica ctu proceedings 24:44–49, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app analysis of neutron fields generated in spallation targets of b-uran experimental assembly using monte carlo method ondřej šťastnýa, b, ∗, miroslav zemana, d, dušan krála, karel katovskýa, elmira melyana, c, jindřich adama, d, alexander alexandrovich solnyshkind a brno university of technology, faculty of electrical engineering and communication, department of electrical power engineering, technická 3082/12, 616 00 brno, czech republic b state office for nuclear safety, department of non-proliferation, senovážné nám. 9, 110 00 prague 1, czech republic c yerevan state university, faculty of physics, 1 alex manoogian street, 0025 yerevan, armenia d joint institute for nuclear research, laboratory of nuclear problems, joliot-curie 6, 141980 dubna, russia ∗ corresponding author: xstast30@stud.feec.vutbr.cz abstract. the aim of this paper is to introduce experimental assembly b-uran and the results of monte carlo simulations of neutron fields, which will be generated by using various spallation targets. this experimental assembly was constructed in joint institute of nuclear research in dubna, russian federation, in order to study accelerator driven systems fundamental characteristics. beam of 660 mev protons should be used for that purpose. the mcnp model of such set-up has been developed at brno university of technology, czech republic. the goal is to get data needed for prediction of reaction rates in detectors placed in b-uran experimental channels. such data will be experimentally validated later. furthermore, simulations of radiation exposure around this experimental assembly were performed. keywords: accelerator driven systems, b-uran, monte carlo simulations, reaction rates, spallation reactions. 1. introduction there is an area of nuclear power production, which is still considered as its weakness – spent nuclear fuel, which contains a lot of highly radioactive fission products and long-lived actinides; therefore, it has to stay separated from the environment forever. deep geological repositories can serve as the most conservative solution for this purpose. nevertheless, there are other options, such as reprocessing, which provides possibility to use plutonium again as a part of mox fuel. however, fission products and long-lived actinides remain as a heritage of nuclear power production and they are treated as radioactive waste. several decades ago, one promising solution was proposed – accelerator driven systems (ads) [1], which have the potential to "incinerate" such kind of problematic isotopes and therefore minimize the amount of radioactive waste after reprocessing the spent nuclear fuel. ads usually consist of accelerator, spallation target and blanket. accelerator produces accelerated particles (protons, deuterons) to high energies (hundreds of mev or ones of gev) which than trigger spallation reactions in target material. spallation reactions produce high-energy neutron fields around the target that can be used to transmute ("incinerate") radionuclides located in surrounding blanket (see the example at figure 1, neutron yield by using 800 mev protons and tungsten target). there has been a lot of research activities worldwide focused on this area so far, but still the compact models and nuclear data libraries describing the parameters of generated neutron fields are missing (for more information see [3–6]). one of the research facilities with long experience in this research is joint institute of nuclear research (jinr) in dubna, russian federation. group e+t raw (energy & transmutation of radioactive waste) with experimental assembly quinta [3] can be mentioned in this regard. local accelerator phasotron produces particles with energies up to 660 mev and various targets are available, as well. this creates a great opportunity for many researches for studying spallation reactions and parameters of generated neutron fields. 2. description of proposed experiment with b-uran experimental assembly 2.1. b-uran experimental assembly latest experimental assembly b-uran was designed at jinr for studies of spallation neutrons behavior in quasi-infinite geometry. this experimental assembly is a cylinder made from depleted uranium with diameter 44 http://dx.doi.org/10.14311/app.2019.24.0044 http://ojs.cvut.cz/ojs/index.php/app vol. 24/2019 analysis of neutron fields generated in spallation targets figure 1. neutron yield from fission and spallation inside tungsten target [2] figure 2. schematic drawings of b-uran, front and rear view [7] of 120 cm and length of 100 cm. it is coated with 10 cm thick steel layer and weights app. 20 tons. central zone allows positioning replaceable cylindrical targets with 20 cm in diameter. for experimental purposes, b-uran offers large amount of axial experimental channels [3]. more details are shown in figure 2. 2.2. spallation targets several types of spallation target were considered – lead, graphite and bismuth. lead target was 100 cm in length, but due to the high leakage of neutrons from the front part of b-uran, it was decided to make a hole 8 cm in diameter and 40 cm in length so the particle beam could penetrate deeper into the target geometry in order to have more convenient distribution of the generated neutron field inside the inner figure 3. b-uran sectional view, lead target [8] part of b-uran. for further details, see figure 3. 2.3. mcnp model of the proposed experiment first, the mcnp 6.2 [9] model of b-uran experimental assembly was developed. it was decided to simplify it in order to decrease the cpu time needed, therefore, only one plane of experimental channels (9 in total) was considered. each experimental channel is 3 cm in diameter, and they are located at positions 14, 18, 22, 26, 30, 34, 38, 44 and 52 cm from the b-uran axis. the source of particles is a proton beam of en45 o. šťastný, m. zeman, d. král et al. acta polytechnica ctu proceedings figure 4. b-uran sectional view, graphite target, red arrow is a symbol of proton beam (created by mcnp) ergy 660 mev impacting on target in b-uran central zone. for neutron spectra measurement a set of 10 threshold activation detectors is placed inside each considered experimental channel along the b-uran’s length. preliminary, it was decided that small "coin" detectors (diameter 2 cm, height 0.5 cm) made from bismuth will be used. the aim is to detect (n,xn) reaction rates up to several hundreds of mev. according to the results of calculation, different materials can be used for this purpose in further research. see the figure 4 for detailed image of b-uran experimental assembly and location of threshold activation detectors inside experimental channels. it should be stated, that the b-uran experimental assembly is located at story with height of 250 cm, so there is 45 cm of space between the ceiling and the top of concrete shielding. both ceiling and the floor are made from 1 cm thick steel plate. all these details correspond with the real conditions at jinr. surrounding environment of b-uran was represented by air in shape of cylinder with radius 200 cm. there were also additional "control volumes" filled with air around b-uran which were used for monitoring of neutron leakage from the experimental assembly. graphite target was 100 cm in length, despite bismuth target was 50 cm only and was placed in the rear position inside b-uran target channel – see figure 4. for neutron and gamma dose calculation, a simplified model of b-uran has been created. it also involved its surroundings (concrete shielding) at the story, where this experimental assembly is located. all calculations were performed using default physics settings of mcnp (model cem03.03 [10]) and photofission settings was turned on. concerning used nuclear data libraries, see table 1. for thermal treating a free-gas model was involved, no s(a,b) lifigure 5. b-uran sectional view, bismuth target, red arrow is a symbol of proton beam (created by mcnp) braries were used. for all other reactions, internal default mcnp models were used. 2.4. expected model outcomes the main aim of the model was to determine neutron flux values inside the threshold activation detectors (tally f4 – flux) in order to get data for further reaction rates calculation. subsequently, a visual information about several types of particles spatial distribution (mesh tally, cmesh type 1 – flux) was also seek. finally, rough estimation of radiation dose around b-uran during irradiation was investigated (mesh tally, rmesh type 1 – dose), as well. 3. results and discussion results of neutron flux calculation for threshold activation detectors can be found in figure 6. average relative error of values for detectors using spallation target made from graphite, lead and bismuth are 0.17% – 0.6% – 1.08%. it clearly shows that graphite target provides several orders higher values than other target materials (lead, bismuth) – this needs to be evaluated by comparison with experimental results. from the neutron yield point of view, graphite target seems to be the best option. however, other features of target materials should be taken into consideration, such as thermal stability, which could be in favor for heavy metals such as lead. original idea was to perform these calculations for targets made from depleted uranium and thorium as well. however, it turned out, that these heavy metals with fast fission factor need longer computation times at ordinary workstation. furthermore, neutron flux mesh was calculated for several targets – graphite, lead and bismuth and results of the neutron flux from various materials is 46 vol. 24/2019 analysis of neutron fields generated in spallation targets library materials endf/b-7.1 nat. c, 14n, 16o, 55mn, 209bi, 234u, 235u, 238u endl-85 nat. ar endf/b-5 nat. cr, nat. ni, nat. pb endf/b-5 t2 group upgrade [11] nat. fe mcplib-v.04 from endf/b-5 (photon data) nat. c, nat. n, nat. o, nat. ar, nat. cr, 55mn, nat. fe, nat. ni, nat. pb, 209bi, nat. u endf/b-7.0 (photonuclear data) 14n, 55mn, 209bi, 235u, 238u la150 (photonuclear data) 16o endf/b-7.0 (proton data) 14n, 16o, 209bi el03 (electron data) [12] nat. c, nat. n, nat. o, nat. ar, nat. cr, 55mn, nat. fe, nat. ni, nat. pb, 209bi, nat. u table 1. nuclear data libraries for mcnp calculations figure 6. comparison of several targets, neutron flux (neutrons/cm2), 9 sets of detectors in experimental channels (calculated by mcnp) possible to see the in figure 7, where is their comparison. these results fully corresponds with graph in figure 6. since the bismuth target was placed in the rear position of the b-uran target channel, the maximum values of neutron flux were located at the end of each experimental channel. in case of lead target, which fills the whole target channel (with hole in the center), the shape of neutron flux graph almost equals to normal distribution. for graphite target, the maximum of neutron flux in each channel can be found at detectors in the middle of b-uran length. moreover, radiation dose around b-uran during irradiation was computed, as well. drawing of the b-uran surroundings including concrete shielding is shown in figure 8. results can be found in figure 9 – neutron and gamma dose. influence of concrete shielding located in front of b-uran is clearly visible as well as a small tunnel which leads the particle beam to target. the different behavior of neutron and gamma ra47 o. šťastný, m. zeman, d. král et al. acta polytechnica ctu proceedings figure 7. comparison of neutron flux mesh of several targets (created by mcnp) diation is clearly visible. this particular concrete shielding provides very efficient protection against gamma radiation, unlike neutron radiation still creates quite compact neutron field behind the shielding, which decreases by three to five orders of magnitude only, probably because of neutron transport through figure 8. b-uran surroundings including concrete shielding [8] figure 9. neutron and gamma dose around b-uran experimental assembly with lead target, at level of proton beam, irradiated by 4e15 protons (created by mcnp) the space between the top of the shielding and the ceiling. 4. conclusions in this paper, the neutron flux for bismuth threshold activation detectors in experimental channels of b-uran assembly was calculated. these values can be then used for reaction rates determination. furthermore, several targets (graphite, lead and bismuth) were taken into consideration and neutron flux mesh was calculated for them, as well. based on values for neutron flux in threshold acti48 vol. 24/2019 analysis of neutron fields generated in spallation targets vation detectors, it can be stated, that graphite target provides several orders higher values than other target materials – this needs to be validated by comparison with experimental results. other question is graphite thermal stability, which could be in favor for heavy metals such as lead. since the targets made from depleted uranium and thorium were proposed as well, the same calculations should be also performed for them in further research. moreover, geometrical properties of targets should be also considered. the aim should be to minimize neutron leakage from the front and rear part of buran assembly and to maximize neutron field usage inside the threshold activation detectors located in experimental channels. as a result, the aim for further research in this area should be to propose a target design combining light and heavy materials together. based on the gained knowledge, most efficient material composition of detectors should be investigated. for purposes of this article, only single material (bismuth) has been considered. considering neutron and gamma dose calculations, the radiation dose around b-uran has been estimated. according to obtained values, which have been calculated for typical irradiation session at jinr phasotron (4x1015 protons), the doses from gamma and neutrons could be very high (see figure 9). following to that, radiation protection measures have to be clearly defined and regularly verified. moreover, the influence of neutron radiation to the concrete shielding should be investigated (especially with regard to possible concrete activation by neutrons). acknowledgements part of this research work has been carried out in the centre for research and utilization of renewable energy (cvvoze: http://www.cvvoze.cz) and part within energy and transmutation of radioactive waste collaboration (e&t raw: http://et.jinr.ru). authors gratefully acknowledge financial support from the ministry of education, youth and sports of the czech republic under brno university of technology specific research programme (project no. fekt-s-17-4784 "new technologies for sustainable power engineering") and under international research cooperative programme of jinr dubna and the czech republic (projects of joint 3+3 committee no. 11/159/2018 "neutron spectral characterization using spectral shield analysis" and 12/159/2018 "study of neutron flux in spallation targets"). references [1] c. bowman, e. arthur, p. lisowski, et al. nuclear energy generation and waste transmutation using an accelerator-driven intense thermal neutron source. nuclear inst and methods in physics research pp. 336–367, 1992. doi:10.1016/0168-9002(92)90795-6. [2] d. král. investigation of thorium utilization in accelerator driven nuclear reactors. master’s thesis, brno university of technology, brno 2017. [3] w. furman, j. adam, l. závorka, et al. recent results of the study of ads with 500 kg natural uranium target assembly quinta irradiated by deuterons with energies from 1 to 8 gev at jinr nuclotron. proceedings of science 2012. http: //www.sujv.cz/cz/index.php?ns=202&id=1000085. [4] a. koning, d. rochman. modern nuclear data evaluation with the talys code system. nuclear data sheets 113 2012. http://www.sciencedirect. com/science/article/pii/s0090375212000889. [5] d. rochman, a. koning, j. sublet, et al. the tendl library: hope, reality and future. proceedings of the international conference on nuclear data for science and technology 2016. bruges, belgium, https://tendl. web.psi.ch/bib_rochman/tendl.nd2016.pdf. [6] j. thomason. the isis spallation neutron and muon source – the first thirty-three years. nuclear instruments and methods in physics research section a: accelerators, spectrometers, detectors and associated equipment 917:61–67, 2019. issn 0168-9002, doi:10.1016/j.nima.2018.11.129. [7] jinr. energy and transmutation of radioactive waste [online], dubna, moscow region, russian federation, 2019 [quote: 2019-05-26], http://et.jinr.ru/pict/pic_buran.jpg. [8] p. tichý, j. adam, a. baldin, et al. experimental investigation and monte carlo simulations of radionuclide production inside the uranium spallation target quinta irradiated with a 660-mev proton beam. epj web conf 204 04003 2019. doi:10.1051/epjconf/201920404003. [9] c. werner. mcnp users manual – code version 6.2. los alamos national laboratory, report la-ur-17-29981 2017. [10] cem03.03. monte-carlo code system to calculate nuclear reactions in the framework of the improved cascade-exciton model. stepan mashnik and arnold sierk: cem0303 user manual, los alamos national laboratory, report la-ur-12-01364 2012. http://www. oecd-nea.org/tools/abstract/detail/psr-0532/. [11] e. arthur, p. young. evaluated neutron-induced cross sections for 54,56fe to 40 mev. los alamos national laboratory report no la-8626-ms 1980. [12] k. adams. electron upgrade for mcnp4b. los alamos national laboratory internal memorandum, x-5-rn(u)-00-14 2000. 49 http://www.cvvoze.cz http://et.jinr.ru http://dx.doi.org/10.1016/0168-9002(92)90795-6 http://www.sujv.cz/cz/index.php?ns=202&id=1000085 http://www.sujv.cz/cz/index.php?ns=202&id=1000085 http://www.sciencedirect.com/science/article/pii/s0090375212000889 http://www.sciencedirect.com/science/article/pii/s0090375212000889 https://tendl.web.psi.ch/bib_rochman/tendl.nd2016.pdf https://tendl.web.psi.ch/bib_rochman/tendl.nd2016.pdf http://dx.doi.org/10.1016/j.nima.2018.11.129 http://et.jinr.ru/pict/pic_buran.jpg http://dx.doi.org/10.1051/epjconf/201920404003 http://www.oecd-nea.org/tools/abstract/detail/psr-0532/ http://www.oecd-nea.org/tools/abstract/detail/psr-0532/ acta polytechnica ctu proceedings 24:1–6, 2019 1 introduction 2 description of proposed experiment with b-uran experimental assembly 2.1 b-uran experimental assembly 2.2 spallation targets 2.3 mcnp model of the proposed experiment 2.4 expected model outcomes 3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0024 acta polytechnica ctu proceedings 26:24–29, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app on evaluation of the three-dimensional isogeometric beam element edita dvořáková∗, bořek patzák czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: edita.dvorakova@fsv.cvut.cz abstract. the exact description of the arbitrarily curved geometries, including conic sections, is an undeniable advantage of isogeometric analysis (iga) over standard finite element method (fem). with b-spline/nurbs approximation functions used for both geometry and unknown approximations, iga is able to exactly describe beams of various shapes and thus eliminate the geometry approximation errors. moreover, naturally higher continuity than standard c0 can be provided along the entire computational domain. this paper evaluates the performance of the nonlinear spatial bernoulli beam adapted from formulation of bauer et al. [1]. the element formulation is presented and the comparison with standard fem straight beam element and fully three-dimensional analysis is provided. although the element is capable of geometrically nonlinear analysis, only geometrically linear cases are evaluated for the purposes of this study. keywords: bernoulli theory, curved beam element, isogeometric analysis, nurbs. 1. introduction the desire for the automatic connection between computer-aided design (cad) and finite element analysis (fea) has been the crucial impulse for the development of isogeometric analysis [2]. the idea of iga is to use the basis functions used for the geometry description in cad also as the approximation functions for the analysis. this results in the possibility of only one model shared between the design and analysis. isogeometric approach introduces into the analysis some other very important features. one of them is a possibility to exactly model arbitrarily curved geometries, including a conic sections which can be only approximated by standard polynomial basis functions. this makes it very convenient for the use in the analysis of curved beams. the focus of this paper is placed on the geometrically nonlinear three-dimensional bernoulli beam [1, 3]. the beam formulation is briefly introduced and the performance over standard fem approaches is evaluated. 2. nurbs-based analysis the most wide-spread technology in cad industry are nurbs (non-uniform rational b-splines). the nurbs curve is obtained by linear combination of cartesian coordinates of the control points p and nurbs functions rpi c(ξ) = n∑ i=1 r p i (ξ)pi. (1) each nurbs function is generated by weighting bspline functions npi with a given weight wi associated with control point pi r p i (ξ) = ni,p(ξ)wi∑n j=1 nj,p(ξ)wj . (2) b-spline functions form a special subset of nurbs functions (corresponding to constant wi) and are derived recursively starting with a piecewise constant functions ni,0(ξ) = { 1 if ξi ≤ ξ < ξi+1 0 otherwise , (3) ni,p(ξ) = ξ − ξi ξi+p − ξi ni,p−1(ξ) + ξi+p+1 − ξ ξi+p+1 − ξi+1 ni+1,p−1(ξ), (4) where p is the degree of the b-spline function, ξi is the coordinate of the ith-knot and ξ ∈ 〈0, 1〉 is a parametric coordinate. the computational domain in isogeometric analysis (iga) is firstly divided into patches, which are further divided into knotspans (often referred to as elements in iga). understanding of knotspans is similar to elements in standard fem, nevertheless higher continuity up to cp−1 between knotspans can be achieved naturally using nurbs, unlike c0 in standard fem. moreover, the nurbs basis functions are generally non-interpolatory. in the knots (points which are dividing patch into knotspans), the continuity can be locally decreased up to c0 by knot multiplication, 24 https://doi.org/10.14311/app.2020.26.0024 https://ojs.cvut.cz/ojs/index.php/app vol. 26/2020 on evaluation of the three-dimensional isogeometric beam element n4,2 ξ knotspan patch continuity: c0 c0c1c∞ c∞ knotspan p = 2, ξ = {0, 0, 0, 0.5, 1, 1, 1} n1,2 n2,2 n3,2 knots control points p1 p2 p4p3 parametric space physical space x y figure 1. description of b-spline (nurbs) finite element geometry. the geometry is modeled by one patch consisting of two knotspans with quadratic nurbs approximation. which is a standard technique in nurbs technology. this procedure imposes the interpolatory behavior of the functions at the particular point. for better understanding of nurbs see fig. 1, where mapping between real and parametric geometry is illustrated. for more information the reader could refer to [4]. 3. beam element formulation the formulation of the presented three-dimensional beam element [1] is based on bernoulli beam theory and accounts for the geometrically nonlinear behaviour. bernoulli kinematics assume that the orthogonality of the cross-section to the center line is preserved after deformation and that the cross-sectional dimensions remain unchanged. the element has four degrees of freedom in each control point: three global displacements u,v, and w and additional degree of freedom corresponding to rotation around the center line ψ. the rotational degree of freedom enables the element to develop a torsion (warping effects are neglected here) and also to model initially twisted beams. 3.1. geometric description in the following, the standard notation using uppercase and lower-case letters for the undeformed and deformed configuration, respectively, is adapted. the beam formulation (see figure 2) starts from threedimensional approximation reduced using bernoulli kinematic assumptions to the displacements of the center line, which is given by the position vector xc (xc). the position vector of the center line is described as a linear combination of the control points coordinates x̂i and the corresponding basis functions rpi xc = ∑ i r p i x̂ i, (5) xc = ∑ i r p i x̂ i. (6) deformed position vector is given as x̂i = x̂i + ûi, (7) where ûi is a vector of global degrees of freedom (u,v,w). additionally to the center line, the cross-section orientation is described by a moving trihedral given by the base vectors ai (ai) with i ∈{1, 2, 3}. a position vector of an arbitrary point of a beam continuum given by coordinates x (x) can be expressed as x(θ1,θ2,θ3) = xc(θ1) + θ2a2(θ1) + θ3a3(θ1), (8) x(θ1,θ2,θ3) = xc(θ1) + θ2a2(θ1) + θ3a3(θ1), (9) where θi are the convective contravariant coordinates. the first components of a moving trihedral a1 and a1 are aligned with a normalized tangents t and t, respectively, and are computed as a1 = ∑ i r p i,1x̂ i, (10) a1 = ∑ i r p i,1x̂ i, (11) where (·),i denotes the derivative with respect to the coordinate θi. the remaining components aα with α ∈{2, 3} are orthogonal to the tangent of the center line (resp. a1) and are described by the reference trihedral, given by a0α, t0, as aα = r̄t(ψ)λ(t0, t)a0α, (12) where ψ is a initial rotation about a tangent of a beam, and λ and r̄t are the two key operations used 25 edita dvořáková, bořek patzák acta polytechnica ctu proceedings a1(θ 1) a2(θ 1) a3(θ 1) a1(θ 1) a3(θ 1) a2(θ 1) u(θ1,θ2,θ3) x(θ1,θ2,θ3) xc(θ 1) xc(θ 1) x y z figure 2. illustration of the beam in its undeformed and deformed configurations. t0 t0 a02 a03 t λ(t0, t ) λ(t0, t )λ(t0, t )a 0 3 a03 λ(t0, t )a 0 2 t0 a02 a03 t λ(t0, t )a 0 3 λ(t0, t )a 0 2 r̄t (ψ) r̄t (ψ) a3 a2 figure 3. the λ(t0, t) and r̄t(ψ) operations used for the alignment of the cross-section at the current position. for the alignment of a moving trihedral illustrated in figure 3. while the mapping matrix λ(t0, t) aligns the reference trihedral given by the tangent t0 to the tangent t at the current position, the rotation matrix r̄t(ψ) rotates the aligned reference trihedral about t with given rotation ψ. the rotation ψ = ψ(θ1) is calculated using basis functions and values ψ̂i assigned to the control points. the euler-rodriguez formula [5] is used to define both the mapping matrix λ and the rotation matrix rt. analogical procedures denoted as λ(t, t) and r̄t(ψ) are used to align the moving trihedral of the undeformed to the deformed configuration. the base vectors aα of the deformed configuration are given as aα = r̄t(ψ)λ(t, t)aα, (13) where ψ = ψ(θi) is a rotational degree of freedom evaluated at the current position θi. the base vectors of the undeformed continuum gi are defined as gi = x,i leading to g1(θ1,θ2,θ3) = a1(θ1) + θ2a2,1(θ1) + θ3a3,1(θ1), g2(θ1) = a2(θ1), (14) g3(θ1) = a3(θ1) with analogical equations for the base vectors gi = x,i of the undeformed configuration. in the sequel, dot products gi · gj ( gi · gj) and ai · aj ( ai · aj) are denoted as gij (gij) and aij (aij), respectively. 3.2. green-lagrange strain tensor the green-lagrange strain tensor e calculated for the curvilinear coordinate system is defined as eij = 1 2 (gij −gij) . (15) and for the orthogonal base vectors, the transformation to the cartesian coordinate system denoted with (̃·) is given by ẽij = eij ‖gi‖2‖gj‖2 . (16) by substituting geometric relations into the strain tensor, the individual components of the tensor can be derived. during the derivation additional simplifications are made assuming that only a slender beam is considered (b,h << l where b and h are cross-sectional dimensions and l is a length of the beam) resulting in the following equations for the diagonal term e11 e11 = 1 2 (a11 −a11)︸ ︷︷ ︸ � +θ2 a2,1 · a1 − a2,1 · a1︸ ︷︷ ︸ κ21 +θ3 a3,1 · a1 − a3,1 · a1︸ ︷︷ ︸ κ31 . (17) the diagonal terms eαα as well as the shear term e23 are equal to zero. this yields from the bernoulli assumptions (undeformable cross–section). for the off-diagonal terms e12,e13 e1α = 1 2 θβ (aβ,1 ·aα − (aβ,1 ·aα)︸ ︷︷ ︸ κβα , (18) where (α,β) ∈{(2, 3), (3, 2)}. 26 vol. 26/2020 on evaluation of the three-dimensional isogeometric beam element 3.3. constitutive equations the energetically conjugated stress tensor to the green-lagrange strain tensor is the second piolakirchhoff tensor s. elastic isotropic material is considered within this work and st. venant–kirchhoff material is applied. as the beam formulation is reduced to the center line and bernoulli theory is assumed, the shear forces s̃23 and s̃32 and normal forces s̃22 s̃33 vanish and the full constitutive relation can be reduced to  s̃11s̃12 s̃13   =   e 0 00 g 0 0 0 g   ︸ ︷︷ ︸ d   ẽ11ẽ12 ẽ13   , (19) where e is young’s modulus, g is the shear modulus and d is a reduced elasticity matrix. 3.4. principle of virtual work the system is in equilibrium when the overall virtual work of the system is equal to zero δw = −δwint + δwext = 0. (20) the internal end external virtual work of the system is given by δwint = ∫ ω s : δedx, (21) δwext = ∫ γn t : δudx + ∫ ω ρ0b : δudx, (22) where ω and γn are the domain and the neumann boundary, respectively. the external virtual work depends on the boundary forces t, body forces b and material density ρ0. by substitution of expressions for the virtual greenlagrange strain tensor obtained from (15), constitutive relations (19) and performing linearization the following incremental relation can be obtained ∑ s krs∆us = −rr, (23) where ∆us are the displacement increments and rr = −(fintr + f ext r ) (24) = ∂wint ∂ur + ∂wext ∂ur , (25) krs = ∂rr ∂us = ∂2wint ∂ur∂us . (26) 4. numerical example the described beam element has been implemented and tested on the example of helicoidal spring (see figure 4). this problem represents fully threedimensional structure where the all membrane, bending and torsion effects are present. the geometry is f = 1 t t e = 108 ν = 0.0 figure 4. helicoidal spring cantilever subjected to the unit force tip load and its cross-section. figure 5. helicoidal spring geometry for different thickness/length ratios: 0.1, 0.05, 0.025, 0.0125. given by xc(θ1) =   10 sin ( θ12π ) 10 cos ( θ12π ) 20 θ1   . (27) the structure is subjected to the unit tip force load. different dimensions of the square cross-section have been used in the analysis, in order to evaluate the performance and the applicability of the underlying assumptions. the studied geometries can be observed in figure 5, where helicoidal springs with different thickness/length (t/l) ratios (0.1, 0.05, 0.025 and 0.0125) are illustrated. the results have been compared with results obtained using standard fem straight beam element with cubic approximation [6]. this element allows for the analysis of both thick and thin beams and is naturally locking-free. the beam has been used in two different configurations: i) the element satisfying bernoulli assumptions; ii) the element accounting also for the shear effects. the results can be observed in the figure 6. as expected, for the thick beam configuration the results obtained using classical fem or iga elements with bernoulli assumptions differ from the standard fem beam element with account for the shear deforma27 edita dvořáková, bořek patzák acta polytechnica ctu proceedings -1.01e-06 -1.00e-06 -9.90e-07 -9.80e-07 -9.70e-07 -9.60e-07 -9.50e-07 -9.40e-07 -9.30e-07 -9.20e-07 -9.10e-07 10 15 20 25 30 35 40 45 t/l = 0.1 -1.60e-05 -1.58e-05 -1.56e-05 -1.54e-05 -1.52e-05 -1.50e-05 -1.48e-05 -1.46e-05 10 15 20 25 30 35 40 45 t/l = 0.05 -2.55e-04 -2.50e-04 -2.45e-04 -2.40e-04 -2.35e-04 -2.30e-04 10 15 20 25 30 35 40 45 t/l = 0.025 -4.05e-03 -4.00e-03 -3.95e-03 -3.90e-03 -3.85e-03 -3.80e-03 -3.75e-03 -3.70e-03 10 15 20 25 30 35 40 45 t/l = 0.0125 iga p=3 iga p=4 iga p=6 fem thin fem shear -1.01e-06 -1.00e-06 -9.90e-07 -9.80e-07 -9.70e-07 -9.60e-07 -9.50e-07 -9.40e-07 -9.30e-07 -9.20e-07 -9.10e-07 10 15 20 25 30 35 40 45 t/l = 0.1 -1.60e-05 -1.58e-05 -1.56e-05 -1.54e-05 -1.52e-05 -1.50e-05 -1.48e-05 -1.46e-05 10 15 20 25 30 35 40 45 t/l = 0.05 -2.55e-04 -2.50e-04 -2.45e-04 -2.40e-04 -2.35e-04 -2.30e-04 10 15 20 25 30 35 40 45 t/l = 0.025 -4.05e-03 -4.00e-03 -3.95e-03 -3.90e-03 -3.85e-03 -3.80e-03 -3.75e-03 -3.70e-03 10 15 20 25 30 35 40 45 t/l = 0.0125 iga p=3 iga p=4 iga p=6 fem thin fem shear -1.01e-06 -1.00e-06 -9.90e-07 -9.80e-07 -9.70e-07 -9.60e-07 -9.50e-07 -9.40e-07 -9.30e-07 -9.20e-07 -9.10e-07 10 15 20 25 30 35 40 45 t/l = 0.1 -1.60e-05 -1.58e-05 -1.56e-05 -1.54e-05 -1.52e-05 -1.50e-05 -1.48e-05 -1.46e-05 10 15 20 25 30 35 40 45 t/l = 0.05 -2.55e-04 -2.50e-04 -2.45e-04 -2.40e-04 -2.35e-04 -2.30e-04 10 15 20 25 30 35 40 45 t/l = 0.025 -4.05e-03 -4.00e-03 -3.95e-03 -3.90e-03 -3.85e-03 -3.80e-03 -3.75e-03 -3.70e-03 10 15 20 25 30 35 40 45 t/l = 0.0125 iga p=3 iga p=4 iga p=6 fem thin fem shear figure 6. comparison of isogeometric beam element with cubic (iga p = 3), quartic (iga p = 4) and hexic (iga p = 6) approximation with standard fem straight beam element with bernoulli assumptions (fem thin) and with account for shear effects (fem shear). the vertical axis corresponds to the vertical tip displacement w, the horizontal axis corresponds to the number of nodes (control points). tion. this difference is diminishing with the smaller thickness/length ratio. in addition to the beam elements, the fully threedimensional analysis has been assessed. the enormous number of nodes, in comparison with beam elements, have to be used in order to obtain sufficiently accurate results. finally, the mesh consisting of 102800 nodes (20×20×257) has been used. for the thin beam with t/l = 0.0125, the obtained deflection is w = −3.877e−3. this result is slightly higher than the results obtained using beam assumptions. the difference between the full 3d and beam models is getting bigger as the height to length ratio increases. this difference is indicating, that the beam theory assumptions are no longer adequate. overall, the results demonstrate the superior convergence properties of the iga element. this is due to its ability to capture curved geometry and natural high-order interpolation. 5. conclusions the spatial isogeometric element based on the work of a. m. bauer et al. [1] has been implemented and its convergence properties have been evaluated in this contribution. the element formulation is geometrically nonlinear, however only the geometrically linear analysis has been considered in this paper. the performance of the presented element has been compared with standard fem beam element using benchmark problem of helicoidal spring. for the higher thickness/length ratios t/l = 0.1, 0.5 the element does not provide good results due to the limitations of bernoulli assumptions. for the thin beams analysed, the excellent results are obtained. acknowledgements the financial support of this research by the grant agency of the czech technical university in prague (sgs project no. sgs19/032/ohk1/1t/11) is gratefully acknowledged. 28 vol. 26/2020 on evaluation of the three-dimensional isogeometric beam element references [1] a. m. bauer, m. breitenberger, b. philipp, et al. nonlinear isogeometric spatial bernoulli beam. computer methods in applied mechanics and engineering 303:101–127, 2016. [2] t. j. r. hughes, j. a. cottrell, y. bazilevs. isogeometric analysis: cad, finite elements, nurbs, exact geometry and mesh refinement. comput methods appl mech engrg 194:4135–4195, 2005. [3] l. greco, m. cuomo. b-spline interpolation of kirchhoff-love space rods. computer methods in applied mechanics and engineering 256:251 – 269, 2013. doi:10.1016/j.cma.2012.11.017. [4] l. piegl, w. tiller. the nurbs book. springer-verlag berlin heidelberg, new york, 1997. [5] j. s. dai. euler–rodrigues formula variations, quaternion conjugation and intrinsic connections. mechanism and machine theory 92:144–152, 2015. [6] z. bittnar, j. šejnoha. numerické metody mechaniky 1. české vysoké učení technické, praha, 1992. 29 https://doi.org/10.1016/j.cma.2012.11.017 acta polytechnica ctu proceedings 26:24–29, 2020 1 introduction 2 nurbs-based analysis 3 beam element formulation 3.1 geometric description 3.2 green-lagrange strain tensor 3.3 constitutive equations 3.4 principle of virtual work 4 numerical example 5 conclusions acknowledgements references doi:10.14311/app.2020.28.0042 acta polytechnica ctu proceedings 28:42–49, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app analysis of fast neutron transport in chloride salts using monte carlo method ondřej šťastnýa, b, ∗, karel katovskýa, dušan krála, antonín krásac, cable kurwitzd, kamil števankaa a brno university of technology, faculty of electrical engineering and communication, department of electrical power engineering, technická 10, 616 00 brno, czech republic b state office for nuclear safety department of non proliferation, senovážné náměstí 9, 110 00 prague, czech republic c belgian nuclear research centre, institute for advanced nuclear systems, sck cen, boeretang 200, 2400 mol, belgium d texas a&m university, department of nuclear engineering, ai engineering building 3133, 423 spence st, college station, tx 77843, united states ∗ corresponding author: xstast30@vutbr.cz abstract. the aim of this paper is to present results of fast neutron behavior analysis within the chloride salts environment using simulations based on monte carlo method (mcnp 6.2). three non-fueled salts (nacl, kcl, mgcl2) and two salts containing fissile material (ucl3, thcl4) were studied. results of this theoretical study will be used for design of an experimental assembly, which will serve to achieve goals of the international research project adar (accelerator driven advanced reactor). one of the project objectives is to investigate chloride salts as potential coolant and a dissolved fuel carrier of advanced nuclear reactor cooled by molten salts and driven by an accelerator. keywords: accelerator driven systems, fast neutrons, adar, chloride salts, mcnp, molten salt reactors, monte carlo simulations. 1. introduction molten salt reactors (msrs) have been identified as promising reactor technology with regard to thorium fuel cycle or for transmutation of spent nuclear fuel from light water reactors. there are several designs based on fuel dissolved in the molten salt coolant, such as fluorides or chlorides [1]. this article is focused on fast neutron behavior analysis within the chloride salts, since they have good nuclear, chemical and physical properties and are much easier to treat. they usually form solutions with lower melting point, and they have higher solubility for actinides. as a result, they can contain significant amounts of transuranic elements. on the other hand, the major isotope of chlorine, cl-35 gets activated to cl-36, long-lived energetic beta emitter, so cl-37 is much preferable in a reactor [1]. as stated above, the goal is to perform studies with fast neutron spectrum since it is used in fast spectrum msrs (msfrs) and or accelerator driven systems [2]. they can have conversion ratios within range burner – converter – breeder. furthermore, they can be optimized for burning minor actinides, for breeding plutonium from u-238, or they may be opencycle power plants since the fast neutron spectrum allows the possibility to omit onsite processing to remove transuranic elements. since the fission products have relatively large neutron capture cross sections in thermal energies, the capture cross sections at higher energies is much lower, resulting in much greater fission product build-up in an msfr than in a thermal-spectrum msr. higher absorption rate at fission products nuclei in thermal-spectrum msr leads to fission products transmutation and therefore they get closer to the valley of stability (after appropriate decay). moreover, the fuel salt heavily loaded with fission products can be sent occasionally for batch processing or allowed to solidify and be disposed of in a repository. for full breeder configuration the fissile material needs to be progressively removed [1]. recent research [3] shows significant discrepancies between evaluated nuclear data libraries in the fast neutron region. improving fast neutron data for chloride salts is essential requirement for the adar project (accelerator driven advanced reactor with molten chloride salt fuel coolant)1 . in order to perform such neutronic research, a theoretical study of an experimental assembly consisting of a fast neutron source in chloride salt environment was performed with the aim to optimize parameters for an experimental campaign using activation detectors. in the following sections, the model is of experimental assembly is described, the current status of 1detailed project information available at https://starfos.tacr.cz/en/project/ltausa18198. 42 http://dx.doi.org/10.14311/app.2020.28.0042 http://ojs.cvut.cz/ojs/index.php/app https://starfos.tacr.cz/en/project/ltausa18198 vol. 28/2020 monte carlo method for fast neutron transport figure 1. salt experimental assembly “solnicka” 3d model relevant nuclear data is summarized and then monte carlo calculations of neutron transport through chloride salts are presented. 2. design of salt experimental assembly 2.1. material in order to study neutronic properties of most commonly available chloride salts, a design of a salt experimental assembly “solnicka” (in english “salt cellar”) has been proposed (fig. 1). the assembly is cylindrical shape with diameter of 50 cm and height of 60 cm with coating from stainless steel (304 type, thickness of 0.6 mm on sides, 1 mm at the bottom and top cover). it contains the following channels (coated with 0.6 mm thick steel layer): • one central channel for neutron source (22 mm in diameter) which extends from top to the middle of the height (30 cm), • four axial experimental channels (28.1 mm in diameter) leading from to the top to the bottom located at radial distance of 5, 10, 15 and 20 cm from the vertical axis, • three radial experimental channels (28.1 mm in diameter) at heights of 35, 45 and 55 cm above the bottom. their length is 23 cm spanning from the outer surface of the assembly in direction to the vertical axis. the experimental channels have been filled with vacuum for simplicity and in order to avoid influence of other materials. as a source of neutrons, the ambe source (cylinder with diameter of 19 mm and height of 40 mm) located in the middle height (30 cm) with emission rate of 5.5615 × 106 neutrons/s was considered since it emits fast neutron field (0.1-11 mev, see table 1 for probability distribution function (pdf) e [mev] pdf e [mev] pdf 4,14e-07 0,00e+00 5,68e+00 2,33e-02 1,10e-01 1,44e-02 5,89e+00 2,06e-02 3,30e-01 3,34e-02 6,11e+00 1,82e-02 5,40e-01 3,13e-02 6,32e+00 1,77e-02 7,50e-01 2,81e-02 6,54e+00 2,04e-02 9,70e-01 2,50e-02 6,75e+00 1,83e-02 1,18e+00 2,14e-02 6,96e+00 1,63e-02 1,40e+00 1,98e-02 7,18e+00 1,68e-02 1,61e+00 1,75e-02 7,39e+00 1,68e-02 1,82e+00 1,93e-02 7,61e+00 1,88e-02 2,04e+00 2,23e-02 7,82e+00 1,84e-02 2,25e+00 2,15e-02 8,03e+00 1,69e-02 2,47e+00 2,25e-02 8,25e+00 1,44e-02 2,68e+00 2,28e-02 8,46e+00 9,68e-03 2,90e+00 2,95e-02 8,68e+00 6,52e-03 3,11e+00 3,56e-02 8,89e+00 4,26e-03 3,32e+00 3,69e-02 9,11e+00 3,67e-03 3,54e+00 3,46e-02 9,32e+00 3,81e-03 3,75e+00 3,07e-02 9,53e+00 5,06e-03 3,97e+00 3,00e-02 9,75e+00 6,25e-03 4,18e+00 2,69e-02 9,96e+00 5,52e-03 4,31e+00 2,86e-02 1,02e+01 4,68e-03 4,61e+00 3,18e-02 1,04e+01 3,70e-03 4,82e+00 3,07e-02 1,06e+01 2,78e-03 5,04e+00 3,33e-02 1,08e+01 1,51e-03 5,25e+00 3,04e-02 1,10e+01 3,63e-04 5,47e+00 2,74e-02 1,11e+01 0,00e+00 table 1. energy spectra for considered ambe neutron source according to iso-8529-1 [5] salt density ! g/cm3 " kef f pucl3 5.71 1.27283 +/0.00006 ucl3 5.50 0.12304 +/0.00001 thcl4 4.59 0.01936 +/0.00000 nacl 2.18 kcl 1.98 mgcl2 2.32 table 2. results of multiplication coefficient calculation (using 250 cycles with 106 particles from the source) overview) standardized iso-8529-1 [5] and can suitably serve for purposes of this study. to reach maximal homogeneity of the salt environment around the neutron source, a salt plug is considered to be placed above the source inside the source channel. for purposes of this study, a model of salt experimental assembly has been created using mcnp 6.2 (see fig. 2) [6]. three non-fueled salts (nacl, kcl, mgcl2) and two salts containing fissile material (ucl3, thcl4) were studied using calculation of reaction rates for set of selected reactions (see fig. 3). the non-fueled salts serve to investigate their potential influence on transport of neutrons (as a potential ternary salt coolant for nuclear reactors [7], [8], [9]). 43 o. šťastný, k. katovský, d. král et al. acta polytechnica ctu proceedings figure 2. salt experimental assembly "solnicka" vertical and horizontal cross section at height of 35 cm (from the mcnp model); nos. 1-4 are assigned to axial channels, 5-7 to radial channels figure 3. cross sections of activation reactions [4] ucl3 (with depleted uranium) and thcl4 represent a chloride salt carrying nuclear material – as a reference chemical compound which could be used in a molten salt reactor, where the fuel is dissolved in the coolant. pucl3 salt has also been considered, but salt assembly with abovementioned dimensions would be supercritical (see tab. 2). activation foils were selected as neutron detectors. the chosen foil materials are ni, in and au. they are mainly mono-isotopic or with one strongly dominating isotope, therefore, the identification of the reaction channel from the observed product is expected to be unambiguous. non-threshold and threshold reactions of these materials are shown in fig. 3. thanks to distribution of the threshold energies between 0.7 and 11.0 mev it will be possible to investigate sensitivity of activation to various parts of the fast component of the neutron field and their changes in the salt environment. 3. results and discussion source driven calculations were carried out with mcnp 6.2 and using the endf/b-vii.0 evaluated nuclear data library. radial and axial reaction rates from the fig. 3 have been calculated in all experimental channels (numbers 1-4 are vertically sorted by increasing distance from central source channel, 5-7 are horizontally sorted by increasing height) with 1010 neutrons from the source (resulting in relative error of results lower than 0.01), see examples of the gold foil activation in the following figures. the observed trends are similar for different activation foil materials (differences are mainly in the absolute values). the results are presented in a systematic comparison in three groups: • one salt, one foil material, one reaction – different radial and axial channels (fig. 4); • one salt, one foil material, one radial or axial channel – different reactions (fig. 5 and fig. 6); • one radial or axial channel, one foil material, one reaction – different salts (fig. 7 and fig. 8) the axial as well as radial distributions of nonthreshold (n,γ) reactions were found to be rather broad compared to the spatial distributions of threshold reactions, which are narrower, likely due to their higher sensitivity to the fast neutrons with energies greater than 1 mev that dominate in the vicinity of the ambe source. in the axial channel closest to the source, the peak factor of the threshold reactions is about an order of magnitude bigger than that of the non-threshold reactions. in the contrary, the peak 44 vol. 28/2020 monte carlo method for fast neutron transport figure 4. axial (top) and radial (bottom) reaction rates of non-threshold (left) and threshold (right) activation reactions in gold foils in a ucl3 salt assembly factors of both reaction types are very similar in the furthest axial channel. in absolute terms, the reaction rates of the nonthreshold reactions are about one and two orders of magnitude bigger than those of the threshold reactions at the positions in the middle part of the assembly and on its boundaries, respectively. comparison of the spatial distributions of reaction rates in different salt environments provide an interesting comparison, see fig. 7 and fig. 8. the spatial distributions of non-threshold reactions strongly differ in different salt environments in any of the axial and radial channels. the most intensive low energy neutron field causing the biggest (n,γ) rates was found to be, as expected, in the ucl3 salt, which contributes to the neutron field by fission. somewhat surprisingly, the second most intensive low-energy neutron field was found in nacl, while thcl4 is on the third place, which shows its rather small contribution of the fission neutrons (compare with kef f in tab. 2). the weakest low-energy neutron field is in kcl. the reaction rates in different salts differ up to a factor of 4. the threshold reactions corresponding to the fast neutron field component show completely different trends. in the channel closest to the neutron source the reaction rates are almost the same for any of the salt environments. with increasing distance of axial channels from the source, the threshold reaction rates start differing in various salts. the strongest fast neutron field was found to be in kcl while the weakest one in ucl3 and thcl4. that suggests that the down-scattering and capture prevail over the fission neutron production in these fissile materials. this behaviour is valid for any of the reactions with threshold above 1 mev, while the inelastic scattering in 115in (eth ≈ 0.4 mev) represents a transition region between threshold and non-threshold reactions, see fig. 9. a strong influence of different salt composition can be also seen on specific behaviour in the radial channel furthest from the assembly centre, where the neutron leakage dominates causing a decrease in the threshold reaction rates for kcl, nacl and mgcl2, see fig. 8. however, in ucl3 and thcl4 fission locally prevails and causes the maximum to remain close to the assembly center axis. this will be especially interesting feature for comparison with the experiment. 4. conclusions monte carlo calculations of fast neutron transport through non-fueled salts (nacl, kcl, mgcl2) as well as neutron multiplication in salts containing fissile material (ucl3, thcl4, pucl3) were performed. the 45 o. šťastný, k. katovský, d. král et al. acta polytechnica ctu proceedings figure 5. axial reaction rates in gold foils in a ucl3 salt assembly figure 6. radial reaction rates in gold foils in a ucl3 salt assembly 46 vol. 28/2020 monte carlo method for fast neutron transport figure 7. axial reaction rates of non-threshold (left) and threshold (right) reactions in gold foils in five salt assemblies 47 o. šťastný, k. katovský, d. král et al. acta polytechnica ctu proceedings figure 8. radial reaction rates of threshold reactions in indium (left) and nickel (right) foils in five salt assemblies figure 9. axial reaction rates of threshold reactions in indium (left) and nickel (right) foils in five salt assemblies 48 vol. 28/2020 monte carlo method for fast neutron transport spatial distributions of threshold and non-threshold reactions in ni, in, au activation foils were investigated. all calculations have been so far performed using the endf/b-vii.0 library. based on presented calculation results for nonfueled salts, it can be stated, that mgcl2 or kcl should be more convenient if neutron fast spectrum is desired. on the other hand, nacl shows higher moderation capabilities. concerning salts fueled with u and th, using of thcl4 provides slightly higher yields of non-elastic reactions with higher threshold than the ucl3. acknowledgements authors gratefully acknowledge financial support from the ministry of education, youth and sports of the czech republic under international research cooperative program between texas a&m university, usa and the brno university of technology, czech republic (“adar – accelerator driven advanced reactor”, ltausa18198). references [1] molten salt reactors. world nuclear association [online]. london, united kingdom: world nuclear association, 2020 [cit. 2020-06-15] available at: https: //www.world-nuclear.org/information-library/ current-and-future-generation/ molten-salt-reactors.aspx. [2] p. mcintyre, s. assadi, k. badgley, et al. accelerator-driven subcritical fission in molten salt core: closing the nuclear fuel cycle for green nuclear energy 1525:636–642, 2013. doi:10.1063/1.4802405. [3] j. c. batchelder, s.-a. chong, j. morrell, et al. possible evidence of nonstatistical properties in the 35cl(n, p)35s cross section. phys rev c 99:044612, 2019. doi:10.1103/physrevc.99.044612. [4] m. chadwick, p. obložinský, m. herman, et al. endf/b-vii.0: next generation evaluated nuclear data library for nuclear science and technology. nuclear data sheets 107(12):2931 – 3060, 2006. evaluated nuclear data file endf/b-vii.0, doi:10.1016/j.nds.2006.11.001. [5] international organization for standardization. iso 8529-1:2001 reference neutron radiations — part 1: characteristics and methods of production. [6] c. j. werner. mcnp user’s manual – code version 6.2. los alamos national laboratory, report la-ur-17-29981 (2017). [7] e. shwageraus, p. hejzlar, m. driscoll. liquid salt cooled flexible conversion ratio fast reactor: neutronic design. nuclear engineering and design 239:2626– 2645, 2009. doi:10.1016/j.nucengdes.2009.07.011. [8] n. e. todreas, p. hejzlar, a. nikiforova, et al. flexible conversion ratio fast reactors: overview. nuclear engineering and design 239(12):2582 – 2595, 2009. flexible conversion fast reactors special section with regular papers, doi:https://doi.org/10.1016/j.nucengdes.2009.07.014. [9] y. li, x. xu, x. wang, et al. survey and evaluation of equations for thermophysical properties of binary/ternary eutectic salts from nacl, kcl, mgcl2, cacl2, zncl2 for heat transfer and thermal storage fluids in csp. solar energy 152:57 – 79, 2017. progress in solar energy special issue: concentrating solar power (csp), doi:10.1016/j.solener.2017.03.019. 49 https://www.world-nuclear.org/information-library/current-and-future-generation/molten-salt-reactors.aspx. https://doi.org/10.1063/1.4802405 https://doi.org/10.1103/physrevc.99.044612 https://doi.org/10.1016/j.nds.2006.11.001 https://doi.org/10.1016/j.nucengdes.2009.07.011 https://doi.org/https://doi.org/10.1016/j.nucengdes.2009.07.014 https://doi.org/10.1016/j.solener.2017.03.019 https://doi.org/10.14311/app.2022.33.0238 acta polytechnica ctu proceedings 33:238–244, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague stability of concrete containing blast-furnace slag following exposure to cyclic elevated temperature ahmad iravania, ∗, volkert feldrappeb, andreas ehrenbergb, steffen andersa a university of wuppertal, institute of structural engineering, department construction materials, pauluskirch str. 11, 42285 wuppertal, germany b fehs building material institute, bliersheimer str. 62, 47229 duisburg, germany ∗ corresponding author: a.iravani@uni-wuppertal.de abstract. concrete is widely used in constructions such as industrial floors or airducts in steeland casting industry where it is often exposed to long-term or cyclic elevated temperatures. for these applications, thermal stability of concrete is of vital importance. the strength reduction due to elevated temperatures depends on the temperature level and concrete composition. in this study, the effects of blast-furnace slag cement (cem iii/a) and basaltic aggregates were investigated at temperatures 250◦c to 700 ◦c in comparison to conventional portland cement (cem i) containing quarzitic aggregates. the concretes were cyclically exposed to high temperatures. special attention was paid to mass loss, residual compressive and residual flexural strength depending on type of cement and aggregate as well as the number of thermal cycles. mass loss and strength loss increased with increasing maximum temperature level, as expected. it was generally observed that concretes containing cem iii/a displayed significantly higher residual mechanical properties for almost all temperature levels. concretes containing a combination of cem iii/awith basaltic aggregates showed significantly higher stability at elevated temperatures compared to other concrete mixtures. it is further shown that apart from the maximum temperature the number of thermal cycles is important for the residual mechanical properties. keywords: blast furnace slag, concrete, cyclic elevated temperature, residual compressive strength, residual flexural strength. 1. introduction durability is one of the most important characteristics of concrete constructions, but it can become decisive for elevated temperature environments including cyclic heating and cooling. applications are for instance concrete floors or airducts in steel and casting industry. the use of concretes with higher long-term thermal stability when exposed to continuous cycles of heating and can reduce the cost and down time for repair or replacement. the behaviour of concrete exposed to single hightemperature events such as fire has been extensively studied [1]. however, the stability of concrete subjected to long-term or cyclic exposure to elevated temperatures is still a research topic [2–4]. physical and chemical changes in concrete such as microcracking caused by different thermal extension of aggregates and cement stone or dehydration of calciumsilicate-hydrate phases in cement stone during are irreversible changes. the hot strength of concrete was compared to its residual strength within the course of cooling process and after reaching ambient temperature. results demonstrated significant differences in strength of the hot, cooling and cooled down concrete [5, 6]. frangi et al. [3] reported additional 10% reduction in the compressive strength comparing residual compressive strength and hot compressive strength, as long as temperature remains below 600 ◦c [3]. lucio-matin et al. [7] report an increasing number of micro-cracks up to 25 cycles when concrete subjected to 1, 25 and 75 cycles between 290 ◦c and 550 ◦c. the cracks initiated in the first heating cycle, with the number of cracks increasing up to the 25th cycle, but remained unchanged henceforth. these findings demonstrate a kind of thermal fatigue in concrete. due to the comparatively low compressive strength of the concrete used here, explosive spalling is not dealt with in this study. it is commonly known that the type of cement, the aggregates as well as the interaction of cement stone and aggregates affect the behaviour of concrete at high temperatures. concretes containing basaltic aggregate or blast-furnace slag as aggregate have shown better thermal stability compared to conventional quarzitic concrete [8–10]. generally, the lower thermal expansion of e.g. basalt or slags reduce thermal stresses and micro-cracking in the concrete structure. as cement paste shrinks during heating, lower thermal expansion of the aggregates is sup238 https://doi.org/10.14311/app.2022.33.0238 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 blast-furnace slag concrete sustainability mix 1 mix 2 mix 3 mix 4 water 286 286 286 286 [kg/m3] cem i 42.5r 550 550 − − [kg/m3] cem iii/a 42.5n − − 550 550 [kg/m3] quartz 0/4 mm 1354 − 1354 − [kg/m3] basalt 0/4 mm − 1534 − 1534 [kg/m3] stabilizer (1% cem.) 5.5 − 5.5 − [kg/m3] w/c-ratio 0.52 0.52 0.52 0.52 [−] initial compressive strength at the age of 46 days 58.4 73.9 65.0 75.4 [mpa] initial flexural strength at the age of 46 days 8.7 11.7 7.0 10.8 [mpa] table 1. overview of concrete mixtures. concrete age climat remarks 1 day − demoulding after casting 1 28 days submerged in water optimal condition for hydration 28 35 days 20 ◦c / 65 % r.h. drying 35 38 days 50 ◦c in dry box reducing water content 38 46 days 20 ◦c / 65 % r.h. hygric homogenization 46 days − initial compressive strength as reference 47 days cyclic temperature exposure starting high temperature exposure temperatures: 250 ◦c, 500 ◦c, 700 ◦c cycles: 1, 5 and 20 table 2. overview of conditioning process. posed to reduce micro-cracking. according to another studies, the residual compressive strength of concrete containing blast-furnace cement (cem iii) is significantly higher compared to concretes with portland cement (cem i), after single temperature exposure. higher residual strengths are reported for concretes containing combinations of cem iii and basalt as aggregate following exposure to cyclic heating [2, 4]. furthermore, alkali-activated binders using ground granulated blast-furnace slags have shown to improve resistance to high temperatures between 500 − 800◦c. this can be attributed to the ability of alkali-activated slags to chemically react forming calcium-silicate-hydrate phases (chs) [10]. therefore, the residual strength of concrete after exposure to cyclic heating/cooling provides essential questions for further investigations concerning the effect of blast-furnace slag in cyclic high-temperature applications. it the aim of this study to quantitatively investigate increases in residual compressive strengths if temperature resistant aggregates, namely basalt, as well optimized binders containing blast furnace slag are used in concrete subjected to thermal cycles at elevated temperatures. as reference concrete conventional concrete made of quarzitic aggregates and ordinary portland cement are tested. apart from the residual compressive strength, residual flexural strength and mass loss are considered. 2. experimental program 2.1. materials and curing of concrete mixes the experimental program is designed to show the effects of blast-furnace slag cement and basaltic aggregates on the residual strengths and mass loss of concrete. as reference concrete made of quarzitic aggregates and ordinary portland cement are used. the resulting matrix of concrete mixtures was fully varied (table 1). in order to be able to compare the results w/cratio and lime content must kept constant. mixture 1 is the reference containing ordinary portland cement (cem i) and quarzitic aggregate. mixture 2 is a combination of cem i and basalt. mixtures 3 and 4 have been developed to investigate the effect of cement type, replacing cem i by blast-furnace slag cement (cem iii/a). specimens were cast in standard prisms (40 × 40 × 160 mm3). the maximum grain size was limited to 4 mm. both cements were supplied by the same manufacturer representing market-based products. thus, the clinker in both cements is the same. 239 a. iravani, v. feldrappe, a. ehrenberg, s. anders acta polytechnica ctu proceedings figure 1. sketch of the heating program, each complete cycle having a duration of 24 h. figure 2. residual compressive strength after 1 heating cycle (a), 5 heating cycles (b) and 20 heating cycles (c). crosses (x) denote untestable mixtures due to complete degradation. in order to reach comparable moisture contents in all mixtures and to reduce the water content to a value usually reached after about 90 days (d) of curing in air, pre-storage was scheduled (table 2). after demoulding, specimens were stored in water for 28 d, followed by 7 d in standard climate (20 ◦c, 60 % relative humidity). afterwards concretes were dried at 50 ◦c for three days in order to reduce the physically bounded moisture. this pre-drying was implemented to reduce the impact of temperature-induced steam pressure build up on the strength of the specimens. finally, the specimens were stored again in standard climate for 7 d to ensure a homogenous moisture distribution. at the time of the first cycle of thermal exposure, specimens had an age of 45 − 48 d. mass loss, residual flexural strength and residual compressive strength were measured after exposure to 1 and 5 heating cycles with the maximum temperature of 250 ◦c, as well as 1, 5 and 20 heating cycles at maximum temperatures of 500 ◦c and 700 ◦c. 2.2. heating programme / temperature regime heating started at ambient temperature and increased with a heating rate of 5 k/min until the desired maximum temperature 250 ◦c, 500 ◦c or 700 ◦c was reached in a conventional tempering furnace. temperature was kept constant for 4 h to ensure the homoňgeneity of the temperature in the specimen and to allow chemical reactions to take place. the cooling process was free but slow cooling for about 16 h until the temperature reached about 50 − 60 ◦c. measurements and strength tests were always carried out after cooling down in ambient condition. figure 2 shows a sketch of the cyclic heating regime. regardless the temperature level, one cycle took 24 h. therefore, specimens which were heated up to 250 ◦c, were stored longer at ambient temperature before the next cycle compared to specimens heated to 700 ◦c. the specimens were examined in terms of water release (mass loss) immediately before testing the residual flexural and residual compressive strength. 3. results and discussion 3.1. residual compressive strength generally, residual compressive strength depends on the maximum temperature during heating. it is demonstrated that the type of cement and aggregate significantly affect the residual compressive strength. furthermore, the number of cycles affect the residual strengths significantly. the residual strengths after heating is expressed relative to the initial strength of concrete prior to thermal treatment at the concrete age of 45 − 48 d, as described in section 2.1. 3.1.1. effect of maximum temperature and number of heating cycles figure 2a (left) displays the residual compressive strength for all mixtures after a single cycle with maximum temperatures of 250 ◦c, 500 ◦c and 700 ◦c. the known increasing reduction of residual strength [1, 5] with an increasing temperature level is confirmed. as expected, basaltic aggregates increase the residual strength compared to quartz at temperature levels of 500 ◦c and higher. 240 vol. 33/2022 blast-furnace slag concrete sustainability figure 3. residual compressive strength for concrete containing cem i and basalt (a) and quartz (b). figure 4. combined effects of type of cement and aggregate on the residual compressive strength after cyclic heating with maximum temperatures of 500 ◦c and 700 ◦c. figure 2b and 2c display the residual compressive strength after 5 and 20 thermal cycles respectively. an ongoing reduction of residual compressive strength can be seen after 5 heating cycles for all tested mixtures (figure 2b). similar to previous findings, the reduction of strength is significant for maximum temperatures of 500 ◦c and 700 ◦c. in addition to the dehydration and degradation of the cement paste, the increase in thermal expansion of aggregates, especially quartz, leads to the formation of micro-cracks resulting in higher degradation. crack formation is further increased by alternating expansion and shrinkage while subjected to thermal cycles. accordingly, stability of concrete significantly depends on the maximum exposed temperature as well as the number of exposed heating/cooling cycles. similar results have been obtained within the current work, which are in accordance with findings e.g. of alonso et al [5]. in figures 3b and 3c it is furthermore obvious that none of the mixtures containing cem i was testable after 20ăthermal cycles due complete degradation to powder. mixtures containing cem iii/a maintained considerable residual compressive strengths even after 20 thermal cycles. 3.1.2. effect of blast-furnace cement and basaltic aggregates the effect of basaltic and quarzitic aggregates on residual compressive strength are depicted in figure 3a and 3.b depending on the maximum temperature. a comparison shows the higher thermal stability of the basalt-containing concretes subjected to 1 and 5 thermal cycles at both temperatures of 500 and 700 ◦c. the residual strength of concrete containing cem i and basalt is about 55% and 30% after exposure to 5 heating cycles at 500 and 700 ◦c, respectively; whereas in the concrete containing quarzitic aggregate, the residual strength is lower than 20% in both cases. this better performance of the basaltic aggregates is attributed to the lower thermal expansion of basalt. nevertheless, both mixtures exhibited complete degradation before 20 thermal cycles were reached. in order to characterize the effects of different types of cements, the residual compressive strengths are compared of mixtures with cem i and cem iii/a are displayed in figure 4. a pronounced decrease in compressive strength is observed for all mixtures before 5 thermal cycles. generally, mixtures containing cem iii/a show higher residual strengths compared to cem i. the stability of cem iii/a containing mixtures is particularly visible after 5 thermal cycles, where cem iii/a obviously reach a maximum 241 a. iravani, v. feldrappe, a. ehrenberg, s. anders acta polytechnica ctu proceedings figure 5. residual flexural strength of the concretes after 1 cycle (a), 5 cycles (b) and 20 cycles (c). figure 6. mass loss induced by cyclic heating for all tested concretes. degradation. it has reached a stable structure without further degradation. the authors currently have no a comprehensive explanation at hand. one explanation is that the powdered blast-furnace slag, which is itself stable against temperature as basalt, changes development of cracks and micro-cracks or the thermally induced shrinkage of the cement paste. another approach is a possible re-hydration tendency of cem iii/a due to the blast-furnace slag and moisture uptake in the air between temperature cycles. this re-hydration could result in a chemical closure of micro-cracks [11, 12]. 3.2. residual flexural strength apart from the residual compressive strength, the residual flexural strength was measured for all mixtures, temperature levels and numbers of cycles. figure 6 displays the results after exposure to 1 (figure 6.a), 5 (figure 6.b) and 20 (figure 6.c) thermal cycles. a significant reduction of the residual flexural strength occurs after any thermal treatment. the reduction is determined both by the temperature level as well as the number of thermal cycles. the loss of residual flexural strengths is generally higher compared to the reduction of the compressive strength. figure 5.b and 5.c show significant stability of concretes containing cem iii/a after 5 and 20 cycles. it is worth mentioning that compared to mixtures with quartz, flexural strengths of mixtures with basalt are higher both, before (see table 1) and after heating to temperatures 500 ◦c and 700 ◦c. this can be explained by to better mechanical properties of basalt. 3.3. mass loss after temperature exposure figure 6 depicts the percentage of mass loss in all concrete mixtures and temperature regimes. the mass loss increases with the maximum temperature. the increase of mass loss is attributed first to the release of 242 vol. 33/2022 blast-furnace slag concrete sustainability physically bound water, followed by beginning of dehydration of csh-phases and the subsequent release of further chemically bound water at higher temperatures. this corresponds to the findings presented in section 3.1.1, where the maximum temperature significantly affected the residual compressive strength for all mixtures. as observed, and described in [12], mass does not change significantly after the first cycle and up to the 5th cycle of heating at temperatures of 500 and 700 ◦c for almost for all mixtures. exposure up to 20 thermal cycles at 500 and 700 ◦c, indicate a slight increase of mass loss for concretes containing cem iii/a. the pronounced mass loss in the first cycle results in shrinkage of the cement stone and subsequent formation of a cracks. this is the reason for high degradation in residual compressive and flexural strengths in the first cycle. this hypothesis is in accordance with findings of lucio-martin et al. [7]. the effect of cement type on the concrete mass loss is evident from figures 6b and 6d, where the higher mass loss for mixtures containing cem iii/a compared to cem i is observed. this might further imply that the bonding of water to the cement paste differs with the type of cement. from the results presented here it could be expected, that water is easier releasable in concrete containing cem iii/a cements. 4. conclusions from the generally known effects of an increasing reduction in residual strengths with increasing temperature, the following conclusions can be drawn: • the residual compressive and residual flexural strength decrease strongest in the first thermal cycle, especially at temperatures of 500 ◦c and higher. the further decrease depends on the type of cement and type of aggregate. the performance gets better if either cem iii/a is chosen as cement or basalt as aggregate. the combination of both shows best results in terms of residual strengths. • at higher temperature levels, concretes containing basaltic aggregates perform better than quarzitic aggregates. • blast-furnace slag cement (cem iii/a) can significantly improve the residual compressive and flexural strengths, particularly after exposure to 20 thermal cycles and in combination with thermally stable aggregates such as basalt. • the mass loss increases with an increasing maximum temperatures, but is not significantly impacted by the number of thermal cycles. it was furthermore observed that the mass loss of cem iii/a is generally about 1 % by mass higher compared to cem i cement. 4.1. residual flexural strength apart from the residual compressive strength, the residual flexural strength was measured for all mixtures, temperature levels and numbers of cycles. figur 5 displays the results after exposure to 1 (figure 5a), 5 (figure 5b) and 20 (figure 5c) thermal cycles. acknowledgements the igf project 20318n of the research association "vdeh-gesellschaft zur förderung der eisenforschung mbh" is funded by the aif within the programs for sponsorship by industrial joint research (igf) of the federal ministry of economic affairs and energy (bmwi) based on an enactment of the german bundestag. we further acknowledge our industrial partners for their project support. references [1] u. schneider. verhalten von beton bei hohen temperaturen, technische universitaät braunschweig, 1982.https://publikationsserver.tu-braunschwei g.de/servlets/mcrfilenodeservlet/dbbs_derivate _00043616/3444-1790.pdf. [2] s. anders. comparison of residual strengths of concretes with quarzitic, limestone and slag sand constituents after cyclic high-temperature exposure. concrete repair, rehabilitation and retrofitting, crc press, p. 223-4, 2015. [3] a. frangi, c. tesar, m. fontana. tragwiderstand von betonbauteilen nach dem brand. bauphysik 28(3):17083, 2006. https://doi.org/10.1002/bapi.200610017. [4] h. y. wang. the effects of elevated temperature on cement paste containing ggbfs. cement and concrete composites 30(10):992-9, 2008. https: //doi.org/10.1016/j.cemconcomp.2007.12.003. [5] k. d. hertz. concrete strength for fire safety design. magazine of concrete research 57(8):445-53, 2005. https://doi.org/10.1680/macr.2005.57.8.445. [6] e. klingsch, a. frangi, m. fontana. experimental analysis of concrete strength at high temperatures and after cooling. acta polytechnica 49(1), 2009. https://doi.org/10.14311/1087. [7] t. lucio-martin t, j. puentes, m. c. alonso. effect of geometry in concrete spalling risk subjected to high temperatures for thermal inertia studies, proceedings of the 6th international workshop on concrete spalling due to fire exposure, p. 71-80, 2019. [8] k. endell. versuche über längen-und gefügeänderung von betonzuschlagstoffen und zementmörteln unter einwirkung von temperaturen bis 1200c: ausgeführt im laboratorium der dozentur für bauwissenschaftliche technologie an der technischen hochschule zu berlin in den jahren 1928/29: ernst; 1929. [9] interner abschlussbericht für die verwendung eines hitzebeständigen betons aus hochofenstückschlacke mit hochofenzement für fundamente eines trockenkühlers internal test report fehs institut für baustoff-forschung (duisburg, germany). 243 https://publikationsserver.tu-braunschweig.de/servlets/mcrfilenodeservlet/dbbs_derivate_00043616/3444-1790.pdf https://doi.org/10.1002/bapi.200610017 https://doi.org/10.1016/j.cemconcomp.2007.12.003 https://doi.org/10.1680/macr.2005.57.8.445 https://doi.org/10.14311/1087 a. iravani, v. feldrappe, a. ehrenberg, s. anders acta polytechnica ctu proceedings [10] interner abschlussbericht für kokerei dünkirchen beton mit hochofenstückschlacke internal test report fehs institut für baustoff-forschung (duisburg, germany). [11] c. alonso, l. fernandez. dehydration and rehydration processes of cement paste exposed to high temperature environments. journal of materials science 39(9):3015-24, 2004. https: //doi.org/10.1023/b:jmsc.0000025827.65956.18. [12] a. iravani and s. anders, effects of cement type, aggregate type and concrete age on the mass loss of concrete exposed to elevated temperature, 5th international workshop on concrete spalling due to fire exposure in, pp.63-72, 2017. 244 https://doi.org/10.1023/b:jmsc.0000025827.65956.18 acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0081 acta polytechnica ctu proceedings 26:81–85, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app the influence of the use of recycled gypsum on the properties of gypsum products in labor conditions hana sekavováa, b, ∗, jakub herrmanna, zdeněk prošeka, miroslav nyčb, george karra’ac a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b knauf praha spol. s r.o., mladoboleslavská 949, 197 00 prague 9, czech republic c lavaris s.r.o., areál šroubáren 43, 252 66 libčice nad vltavou, czech republic ∗ corresponding author: hana.sekavova@fsv.cvut.cz abstract. the article presents the first results from research, which is deal with recycling of plaster boards and use of the resulting material. the research is carried out within the project mpo trio č. fv30359 “recyklace sádrokartonových desek a nová materiálová využití s přidanou hodnotou – gipsrec”. the plasterboards are produced of stucco, which is calcinated of flue gas desulfurization gypsum (fgd gypsum). there is possibility to replace a part of this stucco with recycled material. there is opportunity for saving costs for production. the question how this utilization can change the properties of the products is very important and it is necessary to find the possible improvement or deterioration. this is the objective of this research. the effect on properties was monitored on the samples which was prepared of only pure stucco and of stucco with part of recycled material. there was observed water ratio (splash test), setting times, volume changes, flexural strength, compressive strength and elasticity module. keywords: fgd gypsum, stucco, recyclation. 1. introduction the gypsum was used as a building material for thousands of years. there are references, that the gypsum occurs on buildings dated 7000 bc. the gypsum was used in ancient egypt and rome. great fame has gypsum in 16th century and it is very popular material till today [1]. there are two main sources of gypsum. at first we can find several sources of natural gypsum in the world. there is only one in czech republic and it is situated near the opava. this is part of polish gypsum pan. the purity of the natural gypsum is variable and it depends on the location of mine. the purity of gypsum in the mine kobeřice u opavy is 60–85 % [1]. higher level of purity has gypsum from industrial operations. the syntetic gypsum is still the waste product of the other production. there are for example fgd gypsum, and gypsum generated during the chemical production [1]. the amount of the natural gypsum is limited and the amount of fgd gypsum depends on working power plants which used coal for producing of electric energy (fgd gypsum is product of wet limestone scrubbing of flue gas contains sulfur oxides chemical reaction between limestone suspension and sulfur oxides). the plan for the future is that these power plants will end their life about year 2050, because of their service life and because of the end of coal mining [2]. it is necessary to search for the new possibilities, how to reduce the amount of gypsum, which is used for production of building material. the first way is to find the fillers and additives, which reduce the amount of stucco in production. and the second way is recycled the waste of the gypsum based products and put them back to the production. there is also the pressure to develop the no-waste-production. the recycling and way to the no-waste-production is very popular and the research how to recycle and reuse old building materials or waste from production is running. very popular is recycling of concrete, or using of recycled materials for concrete as a filler. the greater interest in this material is due to the large use of concrete in construction. the observation of the influence on mechanical properties in this area can be found in many researches by j. topič, p. tesárek et al. [3], [4], [5]. there is not only concrete in construction and for example gypsum based building materials are very popular for their properties and it is necessary to deal with this issue. gypsum products have a relatively low bulk density, which ranges from 600–1200 kg m−3, which is about half the bulk density of conventional concrete. this feature allows easy handling of products during construction and the another advantage is less load on the building structure. gypsum is used for its properties in fire protection construction. the gypsum looses during effect of high temperature crystal bound water. this water makes the fire protection of the construction, because 81 https://doi.org/10.14311/app.2020.26.0081 https://ojs.cvut.cz/ojs/index.php/app h. sekavová, j. herrmann, z. prošek et al. acta polytechnica ctu proceedings it decreases the temperature of the surface of the construction [6]. the another positive property of gypsum is that it is hygroscopic. it means, the rehydrated gypsum can absorb the air humidity. it is the regulator of humidity in living space, what brings comfort feeling in rooms. the volume changes are relative low during this process [7]. 2. material the materials, which were used for these experiments, were the white stucco and the waste material from two sources. the white stucco is calcinated of fgd gyspsum. the fgd gypsum is produced during the desulphurisation process flue gases in power plant, where is burned the brown coal. this fgd gypsum has very high level of purity (96 % on average). the white stucco calcinated of fgd gypsum is used as main material in production of plaster boards and gypsum plasters. these materials are showed on figure 1. the first option of waste material were the gypsum based blocks. the second option of waste material were the plaster boards (pb). both of these waste materials were crushed and milled. in the case of boards there was separated the paper above that and the grains were less fine than the grains of milled gypsum blocks. the crushing, milling and separating was done on lavaris equipment. figure 3 shows the material after primer crusher. this material included both of main parts of plaster boards – gypsum core and paper. figure 4 shows material without paper. the paper was separated and the gypsum material is ready for sorting. the fine milled material and paper are showed on figure 2. figure 1. natural and fgd gypsum, stucco. figure 2. fine milled gypsum and separated paper. figure 3. crushed plaster boards with paper. figure 4. milled gypsum before sorting and without paper. 82 vol. 26/2020 reuse of recycled gypsum amount of stucco amount of recycled material initial time final time splash [% wt.] [% wt.] [min] [min] [mm] 90 10 1 2 50 75 25 2 2 56 50 50 12 > 2 weeks 58 10 90 > 2 weeks > 2 weeks 61 table 1. basic properties. figure 5. volume changes over time – recycled blocks [8]. 2.1. the effect of recycled material on basic properties in the first step was observed the behaviour the gypsum slurry depending on the amount of recycled material in recipe. the slurry was mixed by hand without any equipment. the water ratio was in each of recipe 1.2. there were observed the setting times (initial time is and final time fs) and the diameter of splash. the results are showed in table 1. these results showed, that the 50 % is the maximal amount of recycled material, which is possible to use for preparing of samples for following tests. the bigger amount leads to loose of solidifying ability. 2.2. the effect of recycled material on volume changes the volume changes was observed on prisms 160 × 40 × 40 mm. there were prepared 3 different recipes for each recycled material, which were selected in first part of experiment. the recipe 1 is reference sample of pure white stucco. the recipe 2 is 90 % of white stucco and 10 % of milled blocks. the recipe 3 is 75 % of white stucco and 25 % of milled blocks. the recipe 4 is 90 % of white stucco and 10 % of milled plaster boards (pb). the recipe 5 is 75 % of white stucco and 25 % of milled plaster boards and the recipe 6 is 50 % of white stucco and 50 % of milled plaster boards. all presented % are % wt. there were prepared 3 samples of each recipe. all of this samples are tested and the average value for each recipe was signed. the prism were measured and weighed over the next 28 days. the samples lost the free moisture (weight loss) and the volume shrinkage could be monitored too. the results are showed in figures 5 and 6. figure 6. volume changes over time – recycled plaster boards [8]. figure 7. volume changes over time – recycled plaster boards [8]. 3. experimental methods and results the shrinkage increased depending on the amount of recycled material. this is because in the samples with more recycled material is more free water, which is not included in crystal bonded water. 3.1. changes of elasticity module over time the prisms prepared for the observed volume changes (recipes 1–3) was used also for observed changes of elasticity module. this property was tested by nondestructive method – modal analysis. the elasticity module was counted on the base of resonant frequency of samples. the measurement was carried out with a impulse hammer with sensor which was linked to software that processed the values of frequencies and after that was possible to count the elasticity module 8. the course of results is showed on figure 7. the elasticity module is increases over time. it depends on 83 h. sekavová, j. herrmann, z. prošek et al. acta polytechnica ctu proceedings recipe content water ratio flexural compressive [nr.] stucco milled blocks milled pb [-] strength strength[% wt.] [% wt.] [% wt.] [mpa] [mpa] 1 100 0 0 1.4 16.85 2.06 2 90 10 0 1.4 11.47 0.63 3 75 25 0 1.4 6.79 0.13 4 90 0 10 1.2 17.55 0.97 5 75 0 25 1.2 21.53 1.48 6 50 0 50 1.2 25.04 1.92 table 2. flexural and compressive strength. lost of free and bounded moisture from the samples. the gypsum core is more compact and stronger. the bigger shrinkage leads to higher values of elasticity module. figure 8. impulse hammer [9]. 3.2. strength destructive methods the prisms were used in the end of this particular research for destructive strength tests. the both tests were executed according to standard čsn 72 2301. the flexural strength is the first part of the test. the test method is a three point flexural test technique and the flexural strength represents the highest stress experienced within the material at its moment of yield 9. the compressive strength test is performed on parts of samples after execution a three point flexural test technique. the force of the machine act at the whole surface of the sample 10. the results are showed in table 2. the decrease of values of compressive strength depending on the increasing amount of milled blocks are completely different from increase of these values depending on the increasing amount of milled plaster boards. the low breaking loads are caused by high content of water in gypsum slurry. the high water ratio was used because of bad consistency of slurry, which was with lower water ratio unprocessable. figure 9. three point flexural test technique [8]. figure 10. compressive strength test [8]. 84 vol. 26/2020 reuse of recycled gypsum 4. conclusions this research found in its very early begins several dependencies in behaviour the gypsum samples which were prepared with a gypsum based recycled materials. that showed that it is very important to find optimal water ratio. optimal consistency of gypsum slurry is necessary to get the gypsum product with the best properties. in practice is usually use the liquifiers, accelerators or retarders to optimize gypsum slurry properties. the optimal water ratio for the gypsum slurry is from 0.6 to 0.8. in these test was necessary to use the water ratio higher, because we used fine milled material which accelerates the solidifying reaction. the fine grain worked as a crystallization nucleus and the reaction was very fast and in this case was reaction slowed down by adding more water. it optimized the process of rehydration, but it brought the decrease the other properties which were tested. they depend on the structure of gypsum core and when the gypsum slurry contents too much water it generates more pores in volume of gypsum based products and this leads to a reduction in strength and bigger volume changes over time. on the other hand it brings also better thermal properties. acknowledgements this paper was based on research supported by grant mpo trio č. fv30359, where cooperate ctu in prague, faculty of civil engineering, lavaris s.r.o. and knauf praha spol. s r.o. references [1] p. rovnaníková, z. keršner. vlastnosti modifikované sádry. first printing. vysoké učení technické v brně, akademické nakladatelství cerm, s.r.o., brno, 2009. [2] v. pačes. budoucnost energetiky očima odborníků. [2019-07-15], https://www.siemens.cz/energo/ budoucnost-energetiky-ocima-odborniku. [3] j. topič, z. prošek. properties and microsture of cement paste including recycled concrete powder. acta polytechnica 57(1):49–57, 2017. [4] p. tesárek, k. šeps, j. topič, et al. influence of recycled materials on resulting mechanical properties of cement composites. modern methods of experimental and computational investigations in area of construction 825:53–56, 2016. doi:10.4028/www.scientific.net/amm.825.53. [5] j. topič, z. prošek, j. hrůza. mechanical properties of cement paste with fine ground recycled concerete depending on the microstructure. modern methods of experimental and computational investigations in area of construction ii 1144:22–27, 2017. doi:10.4028/www.scientific.net/amr.1144.22. [6] a. vimmrová, l.svoboda. stavební hmoty vápenosíranová pojiva. first printing. jaga group, s.r.o., bratislava, 2004. [7] w. schulze. necementové malty a betony. first printing. sntl, praha, 1990. [8] j. herrmann. možnosti recyklace výrobků na bázi sádry, 2019. https://dspace.cvut.cz/bitstream/handle/10467/ 84255/f1-dp-2019-herrmann-jakub-dp_herrmann_ bez%20podpisu.pdf?sequence=-1&isallowed=y. [9] omega-engineering, 2019. https://in.omega.com/pptst/ih101.html. 85 https://www.siemens.cz/energo/budoucnost-energetiky-ocima-odborniku https://www.siemens.cz/energo/budoucnost-energetiky-ocima-odborniku https://doi.org/10.4028/www.scientific.net/amm.825.53 https://doi.org/10.4028/www.scientific.net/amr.1144.22 https://dspace.cvut.cz/bitstream/handle/10467/84255/f1-dp-2019-herrmann-jakub-dp_herrmann_bez%20podpisu.pdf?sequence=-1&isallowed=y https://dspace.cvut.cz/bitstream/handle/10467/84255/f1-dp-2019-herrmann-jakub-dp_herrmann_bez%20podpisu.pdf?sequence=-1&isallowed=y https://dspace.cvut.cz/bitstream/handle/10467/84255/f1-dp-2019-herrmann-jakub-dp_herrmann_bez%20podpisu.pdf?sequence=-1&isallowed=y https://in.omega.com/pptst/ih101.html acta polytechnica ctu proceedings 26:81–85, 2020 1 introduction 2 material 2.1 the effect of recycled material on basic properties 2.2 the effect of recycled material on volume changes 3 experimental methods and results 3.1 changes of elasticity module over time 3.2 strength destructive methods 4 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0226 acta polytechnica ctu proceedings 33:226–231, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague utilization of recycled and secondary materials in concrete production lca anna horáková∗, alena kohoutková, iva broukalová czech technical university in prague, department of concrete and masonry structures, faculty of civil engineering, thákurova 2077/7, 166 29 praha 6 dejvice, czech republic ∗ corresponding author: anna.horakova@fsv.cvut.cz abstract. the paper describes an assessment of concrete in terms of environmental impacts in relation to the utilization of recycled materials. the article includes a short summary of the literature search on evaluation methods for environmental impacts and on recycled and secondary materials. the environmental impacts of several concrete mixtures were calculated. the reference concrete mixture containing only cement as a binder and only natural aggregate was compared with other mixtures containing recycled and secondary materials as a binder or aggregates. some of these concrete mixtures were then compared within the whole structure. variant design of a simple structure was performed, whereas the variants differed in the concrete mixture. these variants were then evaluated in terms of environmental impacts. keywords: concrete, environmental impacts, life cycle assessment, recycled and secondary materials, variant design. 1. introduction in past years, the issue of sustainable development and the impact of construction activities on the environment are gaining importance. it is desirable to minimize negative environmental impact by suitable design, optimal manufacturing process and material selection. the environmental aspects of sustainable development in the construction industry consist also in the utilization of secondary raw materials in the design and construction of new structures. 1.1. state of the art most often, studies deal with comparison conventional concrete and concrete containing waste or recycled materials such as fly ash, slag or recycled aggregate. the comparison of unit volumes of different concrete mixtures almost always shows, that lower cement dosages make the concrete more environmentally friendly. cement production is responsible for a significant amount of released harmful emissions and its energy consumption is very high. hence, it plays a crucial role in the overall environmental impact of a concrete structure and the cement content in the concrete mixture is a key factor for its environmental assessment. for example, a study [1] reported that a 30% reduction of the cement content leads to a 26.6% reduction in co2 emissions. similar results observed study [2]. the decrease of cement dosage leads also to an energy and raw materials savings. a study [2] reported a 21 % decrease in energy consumption and a 4,3% decrease in raw materials consumption for a 35% reduction of cement content in concrete. the fact, that the environmental impacts of the concrete structure are largely dependent on the cement consumption, led to many research projects dealing with a utilization of latent hydraulic materials as a partial replacement of cement. these supplementary cementitious materials (scm) are capable of hydrating when used together with cement. typical examples are fly ash, blast furnace slag and silica fume. the use of these materials as a partial cement replacement reduces the environmental impacts of concrete production. moreover, the scm are usually waste products, which are generated during the production of other materials or during the energy production. for example, fly ash is a by-product of thermal power plant electricity production and blast furnace slag is a by-product of steel production. thus, there is another environmental advantage, which lies in a utilization of waste products. it should be mentioned that the replacement of cement by scm generally does not leads to a decrease in strength of the concrete. the strength of the concrete containing these materials can be comparable or even higher. cement production harms the environment especially in terms of global warming and climate change and in terms of energy consumption. hence, most studies deal with these environmental impacts and the environmental benefits of using cement replacement materials. the study [3] compares several concrete mixtures with different fly ash replacement levels (25%, 30% and 40%). with the increasing cement replacing level, carbon dioxide emissions decrease. for replacing level of 25%, 30% and 40%, the study reported a decrease in co2 emissions by 6%, 11% and 23%. deterioration of mechanical properties occurred until the replacement level reached 40 %. on the contrary, these mechanical properties improved for lower 226 https://doi.org/10.14311/app.2022.33.0226 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 analysis of concrete and cement epd ref fa bfs wpa wtr rca pc cement [kg/m3] 380 280 324 380 410 380 0 water [kg/m3] 190 236 180 200 200 190 0 fine aggregate [kg/m3] 705 1142 681 572 840 705 1580 coarse aggregate [kg/m3] 1100 493 1160 1020 960 0 0 fly ash [kg/m3] 0 95 0 0 0 0 0 blast furnace slag [kg/m3] 0 0 36 0 0 0 0 waste plastic aggregate [kg/m3] 0 0 0 143 0 0 0 waste tyre rubber [kg/m3] 0 0 0 0 40 0 0 recycled concrete aggregate [kg/m3] 0 0 0 0 0 1100 0 waste pet (as a binder) [kg/m3] 0 0 0 0 0 0 472 superplasticizer [kg/m3] 2 1.8 0 0 0.8 2 0 table 1. the composition of the concrete mixtures. cement replacement levels. for example, at 30 % replacement level, the compressive strength increased by 10 %. if the study was carried out for a real structure, the difference in environmental impact would be more significant than when comparing unit quantities of concrete. this is due to possible reduction of the dimensions of supporting structural elements. nevertheless, it should be mentioned, that the difference would not be significant in case of this study, because there is not a significant increase in strength. some studies deal with the utilization of blast furnace slag as a partial replacement of cement. these studies reported, a reduction of environmental burden, when cement is partially replaced. for example, study [4] reported, that for 50 % replacement level, co2 emissions reduced by 39 %. the study [2] reported similar results. the reduction of co2 emissions, and thus the influence on global warming is the most significant environmental benefit. on the other hand, replacing cement by blast furnace slag has only little effect on raw material consumption. many studies deal with comparison conventional concrete and concrete containing recycled aggregate, which is produced by crushing waste concrete or bricks. according to most of these studies, use of this recycled aggregate leads to reduction of released harmful emissions and consumed raw materials and energy. furthermore, utilization of these waste materials is another benefit. for example, the study [5] compares the environmental impacts of the production of conventional concrete containing only natural aggregate and concrete containing recycled aggregate. the natural aggregate was partially or fully replaced by crushed waste concrete. according to this study, the use of recycled aggregate was advantageous for most environmental impact categories. this was the most significant for the consumption of raw materials, which decreased by 47 % when natural aggregate was fully replaced by recycled aggregate. the reduction of energy consumption was less significant, about 30 % for fully replaced natural aggregate. study [6] reported, that the benefit of using recycled aggregate strongly depend on the transport distance of recycled aggregate. if recycled aggregate is transported over a long distance, the environmental burden caused by the transport may outweigh the environmental benefits of using waste material. 2. methods this paper deals with the utilization of recycled and waste materials in concrete production and its advantages in terms of environmental impacts. at first, the assessment in terms of environmental impacts was performed for concrete, which contains only cement as a binder and only natural aggregate (ref) [7]. then, the assessment was performed for other concrete mixtures, which contain secondary or waste materials as a binder or aggregate. in two of these material variants, there was cement partially replaced by fly ash (fa) [8] or by blast furnace slag (bfs) [9]. in some other variants, there was natural aggregate partially or fully replaced by waste plastics aggregate (wpa) [10], crushed waste tyre rubber (wtr) [11] or recycled concrete aggregate, obtained from demolished buildings (rca) [7]. one more material variant was included in this evaluation innovative material called polymerconcrete, which is composed of aggregate and plastic waste, which replaces cement as a binder (pc) [12]. this material is made from fine aggregate (aggregate size 0 − 4 mm) and waste pet, which is cut into small pieces. the production consists in homogenizing the mixture of plastic waste and small aggregate at high temperature. according to the experiments performed in [13], average compressive strength of this material is 22,7 mpa and average flexural strength is 7,91 mpa. all these mixtures were compared with each other. in the next step, a simple reinforced concrete structure reinforced concrete frame was designed from selected concrete mixtures. then, these structures were compared in terms of environmental impacts. the composition of the concrete mixtures evaluated in this paper is shown in the table 1. 227 a. horáková, a. kohoutková, i. broukalová acta polytechnica ctu proceedings environmental impact explanation global warming global warming is a long-term increase in global average temperaturecaused by excessive production of greenhouse gases. acidification acidification is the ongoing decrease in the ph of the environment. this phenomenon is caused by the presence of acid-forming substances in the atmosphere, which react with water to form acids. eutrophication eutrophication of the environment leads to ecosystem disturbance due toexcessive nutrients in water and soil due to excessive fertilization. photochemical oxidant creation photochemical oxidants are air pollutants that are formed under the influence of sunlight by complex photochemical reactions in air that contain nitrogen oxides and reactive hydrocarbons and cause damage to organisms. abiotic depletion abiotic depletion refers to the depletion of abiotic resources such asfossil fuels, minerals, and clay. table 2. environmental impacts. 2.1. life cycle analysis the assessment of materials and structures in terms of environmental impacts was performed using lifecycle assessment (lca) according to relevant standards [14]. the lca approach is usually based on the whole life-cycle of the investigated product or at least its significant part. so, the assessment includes obtaining raw materials, their transport to the place of processing, manufacturing of the final product, use of the product and further maintenance or repairs if necessary, and final disposal of the product. however, the prediction of the course of the phase of use is sometimes not possible. in these cases, the evaluation includes only a part of the life cycle, it includes for example only obtaining raw materials, their transport to the place of processing and manufacturing of the final product ("cradle to site" evaluation). in this paper, the assessment was performed for concrete as a material for further use, and for a concrete structure too. the evaluation of a unit quantity of different concrete types included the phase of obtaining of raw materials, their transport and processing and manufacturing of the final product. the evaluation of a concrete structure included these phases too, and it includes in addition the phases of transport of concrete and steel and manufacturing of final structure. at first, the consumption of raw materials and emissions released into the environment were defined for each interproduct such as cement, aggregate, water, fly ash, blast furnace slag and other materials. within the assessment, the most significant environmental impacts were considered: consumption of raw materials, global warming and climate change, acidification and eutrophication of the environment and photooxidant formation. in lca, these environmental impacts are called impact categories. principles of these environmental impacts are explained in table 2. the environmental impacts of each interproduct were calculated. impacts on the environment were quantified by so-called impact category indicators measurable variables that can be used to observe changes in the environment. the values indicate the extent of environmental damage caused by human activities. usually, the impact category is influenced by various of substances where some substances are very harmful, and some less. thus, all the substances are converted to an equivalent amount of the reference substance (for example carbon dioxide for global warming and climate change or sulfur dioxide for acidification of the environment). the effect of a specific substance to each impact category was determined by so-called characterization models. a characterization model for a specific impact category is a set of values that reflect the ability of various substances damage the environment within the impact category. all of the issued substances are converted to the equivalent amount of a reference substance by using these values (characterization factors cf). this paper used a characterization model which is recommended in product category rules (pcr) for concrete products [1]. the resultant impact category indicator was calculated according to the following relationship: v xy = cf 1, xy · ! m1i + cf 2, xy · ! m2i + ... + cf n, xy · ! mni (1) where v xy is a result of the impact category indicator (xy indicates the impact category), cf is a characterization factor and m is an amount of a released substance. when evaluating the impact categories, emissions of following substances were considered: carbon dioxide co2, sulfur dioxide so2, nitrogen oxides nox, carbon monoxide co, methane ch4, non-methane volatile organic compound nmvoc, nitrous oxide n2o, hydrochloric acid hcl, hydrofluoric acid hf, hydrogen sulfide h2s, ammonia nh3. in this paper, seven material variants in the amount of 1 m3 were compared (table 1). in addition, several of the above-mentioned material vari228 vol. 33/2022 analysis of concrete and cement epd figure 1. comparison of 1 m3 of concrete mixtures regarding sustainability. ants were compared in designed structure: conventional concrete, concrete with fly ash, concrete with blast furnace slag, concrete with plastic aggregate, concrete with waste tyre rubber and concrete with recycled concrete rubber the structure design was performed for all material variants except polymerconcrete. because of the lack of experience in reinforcing of this material, the load-bearing structure designed from this material was not included in the assessment. a simple reinforced concrete frame was designed from different types of concrete. because of different strength of the concrete types, the volume of concrete needed for the construction depends on the material variant. the amount of steel is the same for all variants. the assessment of the material variants in terms of environmental impacts depends strongly on the volume of concrete, which is used for the structure. therefore, the frame was designed with utilization of the highest load-bearing capacity of the structural elements to make this assessment relevant. the compressive strength of the mixtures and volume of concrete for the variants are shown in table 3. variant the compressive strength[mpa] the volume [m3] ref 32.3 740.00 fa 36.2 728.40 bfs 36.2 728.40 wpa 29.5 749.86 wtr 33.8 735.36 rca 29.2 751.60 table 3. the compressive strength of the concrete and volume ofăconcrete for designed variants. 3. results 3.1. results for 1m3 of concrete the results of the sustainability assessment are related to the specific environmental impact. the following figure shows the comparison of the concrete mixtures for considered impact categories. polymerconcrete was evaluated for all most impact categories as the best material variant. this is probably because this material does not contain any cement, whose production is very burdening for the environment. on the other hand, the energy consumption (the impact category abiotic depletion potential adp fossil) is relatively high, higher consumption was calculated only for conventional concrete. the obvious reason is the high heat consumption for melting the waste polymer. moreover, the use of this material for structures is limited and there is a lack of previous experience with its production. very favourable results have also been obtained for concrete with fly ash. the obvious reason is the reduction of cement consumption due to its partial replacement by fly ash. in the case of concrete with blast furnace slag, cement is also partially replaced, but the replacement level is lower. hence, the results are less favourable for this material variant. material variants, in which recycled materials replaced aggregates, were evaluated as less favourable in terms of environmental impacts. this is because the production of aggregates does not cause such a high environmental burden as the production of cement. another reason is that the waste material (concrete, plastic, tyre rubber) for production of aggregates replacement must be mechanically processed (crushed). this process is energy intensive and reduces the environmental benefits of using waste material. therefore, for reducing the environmental impacts it is preferable to replace cement. 3.2. results for concrete structures for comparison of the whole structures, a simple reinforced concrete frame was designed from different types of concrete. regarding the environmental impacts of concrete frames designed from different material variants, the results did not differ significantly from the previous comparative study (comparison of unit quantities of concrete mixtures). all investigated concrete mixtures have the similar compressive strength. therefore, the dimensions of the supporting elements of the frame do not differ significantly and consumption of concrete is very similar for all designed variants. it is obvious, that a partial replacement of cement by scm is the most advantageous in terms of environmental impacts. the figure 3 shows the comparison of the designed variants for considered impact categories and figure 2 shows the sketch 229 a. horáková, a. kohoutková, i. broukalová acta polytechnica ctu proceedings of the designed frame. figure 2. designed concrete frame. figure 3. comparison of the concrete frames regarding sustainability. 4. conclusion according to this study, in terms of environmental impacts, it is most advantageous to replace a part of cement by supplementary cementitious materials (scm), such as fly ash or blast furnace slag. cement production causes a large environmental burden and many environmental impacts are significantly dependent on cement consumption. it could be reduced by partial replacement of cement by scm. furthermore, supplementary cementitious materials are often waste or secondary materials and their utilization is another environmental benefit. polymerconcrete is also very favourable from the environmental point of view, especially in terms of consumption of raw materials. however, the production of this material is energy intensive, due to the melting of waste plastic. moreover, so far there is no experience with its use in structures. therefore, it is desirable to find a suitable use for this material, for example non-bearing part of structures, such as floors or pavements. when choosing a material for a specific structure, it is necessary to investigate the influence of the transport of recycled materials to the structure site. the transport distance of recycled materials, especially recycled concrete aggregate or waste plastics aggregate, is probably greater than the transport distance of natural aggregates. the utilization of recycled material in concrete usually brings benefits in terms of environmental impacts. on the other hand, the influence of the use of recycled materials on the mechanical properties of concrete must be considered. the use of some waste materials may increase uncertainties in the mechanical behaviour of concrete due to the variability of the properties of these materials. this is typically related to materials that are not commonly used in concrete production, such as recycled concrete aggregate, waste plastics aggregate or waste tyre rubber. hence, it is more appropriate for load-bearing structures to reduce the environmental impacts by using scm. acknowledgements this outcome has been achieved with the financial support of the ministry of education, youth and sports, project: durability of concrete structure and assessment of its life cycle, sgs19/149/ohk1/3t/11. references [1] v. c. li, m. lepech, s. wang, et al. development of green engineered cementitious composites for sustainable infrastructure systems, proceedings of structural engineering world congress (sewc), proc., int’l workshop on sustainable development and concrete technology, beijing, china, iowa state university, p. 181-192, 2004. [2] j. r. prusinski, m. l. marceau, m. g. vangeem. life cycle inventory of slag cement concrete eighth canmet/aci eighth canmet/aci 26 international conference on fly ash, silica fume, slag and natural pozzolans in concrete, 2004. https://www. slagcement.org/portals/11/files/pdf/sus-life-c ycle-inventory-of-slag-cement%20concrete.pdf. [3] s. h. smith, s. a. durham. a cradle to gate lca framework for emissions and energy reduction in concrete pavement mixture design. international journal of sustainable built environment 5(1):23-33, 2016. https://doi.org/10.1016/j.ijsbe.2016.01.001. [4] j.-p. won, h.-h. kim, s.-j. lee, et al. carbon reduction of precast concrete under the marine environment. construction and building materials 74:118-23, 2015. https: //doi.org/10.1016/j.conbuildmat.2014.09.074. [5] n. serres, s. braymand, f. feugeas. environmental evaluation of concrete made from recycled concrete aggregate implementing life cycle assessment. journal of building engineering 5:24-33, 2016. https://doi.org/10.1016/j.jobe.2015.11.004. [6] s. marinković, v. carević. comparative studies of the life cycle analysis between conventional and recycled aggregate concrete. new trends in eco-efficient and recycled concrete, p. 257-91, 2019. https: //doi.org/10.1016/b978-0-08-102480-5.00010-5. [7] k. rahal. mechanical properties of concrete with recycled coarse aggregate. building and environment 42(1):407-15, 2007. https://doi.org/10.1016/j.buildenv.2005.07.033. 230 https://www.slagcement.org/portals/11/files/pdf/sus-life-cycle-inventory-of-slag-cement%20concrete.pdf https://doi.org/10.1016/j.ijsbe.2016.01.001 https://doi.org/10.1016/j.conbuildmat.2014.09.074 https://doi.org/10.1016/j.jobe.2015.11.004 https://doi.org/10.1016/b978-0-08-102480-5.00010-5 https://doi.org/10.1016/j.buildenv.2005.07.033 vol. 33/2022 analysis of concrete and cement epd [8] a. oner, s. akyuz, r. yildiz. an experimental study on strength development of concrete containing fly ash and optimum usage of fly ash in concrete. cement and concrete research 35(6):1165-71, 2005. https: //doi.org/10.1016/j.cemconres.2004.09.031. [9] r. p. salvador, d. a. s. rambo, r. m. bueno, et al. on the use of blast-furnace slag in sprayed concrete applications. construction and building materials 218:543-55, 2019. https: //doi.org/10.1016/j.conbuildmat.2019.05.132. [10] z. z. ismail, e. a. al-hashmi. use of waste plastic in concrete mixture as aggregate replacement. waste management 28(11):2041-7, 2008. https://doi.org/10.1016/j.wasman.2007.08.023. [11] r. chylik, t. trtik, j. fladr, et al. substitution of aeration additive with recycled rubber and its effect on compressive strength and modulus of elasticity. iop conference series: materials science and engineering 583(1), 2019. https://doi.org/10.1088/1757-899x/583/1/012021. [12] a. horáková, j. novák. utilization of recycled plastic for plastic-based concrete. iop conference series: materials science and engineering 596(1), 2019. https://doi.org/10.1088/1757-899x/596/1/012033. [13] a. horáková, j. novák. utilization of recycled plastic for plastic-based concrete. iop conference series: materials science and engineering 596(1), 2019. https://doi.org/10.1088/1757-899x/596/1/012033. [14] wbcsd. technology roadmap: low-carbon transition in the cement industry, 2018. https: //www.wbcsd.org/sector-projects/cement-sustain ability-initiative/resources/technology-roadmap -low-carbon-transition-in-the-cement-industry. 231 https://doi.org/10.1016/j.cemconres.2004.09.031 https://doi.org/10.1016/j.conbuildmat.2019.05.132 https://doi.org/10.1016/j.wasman.2007.08.023 https://doi.org/10.1088/1757-899x/583/1/012021 https://doi.org/10.1088/1757-899x/596/1/012033 https://doi.org/10.1088/1757-899x/596/1/012033 https://www.wbcsd.org/sector-projects/cement-sustainability-initiative/resources/technology-roadmap-low-carbon-transition-in-the-cement-industry https://doi.org/10.14311/app.2022.33.0032 acta polytechnica ctu proceedings 33:32–38, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague stainless steel-frcm system for strengthening of rc beams: towards a sustainable strengthening technique francesco bencardino∗, mattia nisticò department of civil engineering, university of calabria, via p. bucci cubo 39b, 87036 rende, cosenza, italy ∗ corresponding author: francesco.bencardino@unical.it abstract. fabric-reinforced cementitious matrix (frcm) systems for strengthening concrete structures are an alternative to traditional techniques. the frcm system is a composite material consisting of two or more layers of cement-based matrix reinforced with dry fibres in the form of open mesh or fabric. when adhered to concrete structural members the frcm system acts as supplemental external reinforcement. many existing reinforced concrete (rc) members exhibit degradation due to the carbonation of concrete and/or corrosion of internal reinforcing steel bars. these rc members can be strengthened using stainless steel, in strip format (unidirectional fibres), embedded in a cementitious based matrix. the system, named stainless steel-frcm, can be applied according to the externally bonded (eb) technique. in order to reduce times and costs of intervention, number of used materials, as well as the amount of chemical compounds, a novel inhibiting-repairing-strengthening (irs) technique is proposed and experimentally tested. using a suitable matrix (thixotropic mortar with passivation properties) the main three operations of steel bars corrosion inhibition/protection, restoring of deteriorated concrete, and installation of the external strengthening can be carried out in one-step. for evaluating the effectiveness of both new strengthening system and installation technique an extensive experimental investigation was planned and developed. a part of the experimental research includes two groups of three rc beams (3.00 m and 4.80 m long): one strengthened with irs technique, one strengthened with eb technique and one control beam. these beams were tested under monotonic loading. further two beams, one beam for each group, strengthened according to irs technique, were also tested under cyclic loading. the experimental results show the validity of the proposed solution in terms of structural performance and environmental sustainability. keywords: innovative materials, repair, strengthening. 1. introduction the repair and rehabilitation process of reinforced concrete (rc) structures and infrastructures is become a fundamental target in last years. several approaches for strengthening concrete members by composites have been developed and increased significantly as an alternative to traditional methods. combining fibres of different origin with inorganic cementitious matrices new class of composite materials were defined under the acronym fabric-reinforced cementitious matrix (frcm) or similar names as textilereinforced systems. lightweight and high strength of frcm systems fit perfectly in the rehabilitation process. the improvement in terms of performance of strengthened local members with frcm composites was experimentally confirmed by several researches [1–6]. in this context, steel-fabric reinforced cementitious matrix (s-frcm) strengthening system has proven to be one of most effective innovative reinforcement used in the civil engineering field [7]. as well as the structural aspect, the use of environmentally sustainable repair and reinforcement techniques and materials is becoming increasingly important. so eco-friendly geopolymeric matrices act as corrosion inhibitor and together to the repair of the cover concrete with inside the steel strengthening fibre represent a new effective application [8–10]. this latter technique, alternative to the traditional externally bonded (eb) application, is proposed with the name of irs (inhibiting/repairing/strengthening) because of it includes these three operations in one. the effectiveness of this methodology of strengthening has been experimentally evaluated on the basis of obtained test results. the irs strengthening procedure is a novel ecofriendly intervention and can be representing the best choice when considering the structural effectiveness as well as sustainable development. in fact, the use of this technique does not require passivating materials of the internal reinforcements (applying geopolymeric matrices), shorter execution times of the intervention (strengthening and restoration/repair of the surfaces in one step), reduced intervention times and consequently less use of skilled manpower. the economic 32 https://doi.org/10.14311/app.2022.33.0032 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 stainless steel-frcm system for rc beams � � � � ����������������� � ���������������� � d�� e� )ljxuh����3uhsdudwlrq�ri�wkh�vxuidfhv�dqg�furvv�vhfwlrq�diwhu�wkh�vwuhqjwkhqlqj�surfhvv��d��(%�)5&0� figure 1. preparation of the surfaces and cross-section after the strengthening process: a) eb-frcm system and b) irs-frcm system. advantage therefore represents a key aspect in the study and development of this irs innovative technique. in order to compare the flexural capacity of beams with the two applications of reinforcement techniques (eb and irs), an experimental flexural test was carried out. the specimens are formed by two groups of beams (a and b) with a length of 3000 and 4500 mm, width of 150 mm for both, 250 and 400 mm of height. the beams were supported by simple supports and a total of eight beams have been tested under four-points bending, subjected to monotonic and cyclic loading. 2. application techniques the preparation of the surfaces and the technique of applying the composite material represents a fundamental phase of the strengthening process. according to the traditional strengthening technique (eb) after the restoration of the cover concrete, if necessary, two layers (of about 5 mm each one) of cementitious matrix are applied. a new innovative application (irs) was experimentally tested with effective results. it consists to restore the deteriorated concrete and to apply external strengthening in one step with installation of the stainless-steel strip in the cover concrete with reduction of time and cost of the intervention. the final layer of mortar shows more consistent thicknesses (30 ÷ 35 mm). details of the two techniques eb-frcm and irs-frcm are outlined in figure 5. the beams were prepared with the eb technique through the cleaning of the surfaces from any type of material (dust and other substances) and subjected to light mechanical scarification and small mechanical carvings. with reference to the irs technique, the specimens were cast in the form work without the cover concrete. the idea is to simulate the existing rc elements with damaged cover concrete, which in the strengthening phase is totally removed. the surfaces were wet to avoid absorption of the mixing water of the matrix. in figure 5 the state of the surfaces before and after the strengthening process is reported. 3. experimental programme 3.1. test specimens starting from previous research [1], eight rc beams subjected to four-points bending were tested in the laboratory of the department of civil engineering at university of calabria. the beams were designed to fail in flexure before in shear, using appropriate internal shear reinforcement (stirrups). two types of concrete beams with different amounts of internal steel reinforcement were used. two 16 mm and 33 francesco bencardino, mattia nisticò acta polytechnica ctu proceedings � d�� e�� f�� figure 2. geometry of control and strengthened beams: a) rc control beam; b) eb technique and c) irs technique. group l b h l/h a a/d as a ρs b a′s c stirrups [mm] [mm] [mm] [-] [mm] [-] [-] [%] [-] [mm] m-a 3000 150 250 12 900 4.1 2!12 0.6 2!8 !8/150 mmc-a 2!14 0.9m-b 4800 400 1500 2!16 0.7c-b 2!14 0.5 aas : tension steel reinforcement. bρs = as/ (bd) : internal tension steel ratio. ca′s : compression steel reinforcement. table 1. geometrical properties and internal steel reinforcement of tested beams. 12 mm steel bars were used as bottom reinforcement in the monotonic tested beam (group a and b, respectively) and 14 mm in the cyclic tested specimens. the shear reinforcement consisted of 8 mm steel stirrups placed at intervals of 150 mm and two 8 mm steel bars as top reinforcement to hold the stirrups for all beams. the aim is to investigate the structural performance of the eb-s/frcm and irs-s/frcm systems considering two large-scale rc beam groups (3000 mm group a and 4800 mm group b). some geometrical dimensions of the two groups have been changed in order to evaluate potential scale effect. beams referred to as m were tested under static monotonic loading; the character c indicates the beams subjected to static cyclic load. to complete the label of the specimens, the abbreviations eb and irs indicate the reinforcement technique used. the internal and external reinforcement ratios and geometry of the beams are presented in figure 2 and table 1. the complete labels of the beams were reported in table 5 (see later). the width (bf ) of s-frcm fabric was equal to 100 mm and 150 mm for the beams of the groups a and b respectively, in order to keep equal the strengthening ratio ρf = (bf tf )/ (bd). where tf is the thickness of the external strip of steel reinforcement. 3.2. material properties and characterization in order to evaluate the mechanical properties of the concrete, compressive and tensile tests were carried out on cylindrical samples (150 mm of diameter × 300 mm of height). the average compressive strength (fcm) at 28 days was 16.8 mpa for group m and 33.5 34 vol. 33/2022 stainless steel-frcm system for rc beams mechanical property m beams c beams compressive strength, fcm [mpa] 16.8 33.5 tensile strength, fctm [mpa] 1.7 3.7 table 2. mechanical property of concrete. mechanical property m beams c beams diameter ![mm] !8 !12 !16 !8 !14 yield strength, fym [mpa] 543.8 367.1 492.0 510.6 501.2 table 3. mechanical and geometrical property of internal steel reinforcement. stainless-steel fibre geopolymeric matrix nominal thickness, tf [mm] 0.24 compressive strength, fcmm [mpa] ≥ 45 tensile strength, ff [mpa] 1470 flexural tensile strength, ftf m [mpa] ≥ 6 ultimate strain, εf u [%] 2.0 elastic modulus, ef [gpa] 190 table 4. mechanical properties of s-frcm system. mpa for group c; whereas 1.7 mpa and 3.7 mpa the average tensile strength (fctm), respectively. the average yield strength of steel (fvm) of bars !8 (group m), !8 (group c), !12, !14, and !16 mm was 543.8, 510.6, 367.1, 501.2, and 492.0 mpa, respectively. moreover, a summary of used concrete and internal steel properties are given in tables 2 and 3. s-frcm composite material system was used for strengthening the beams. it consists of stainless-steel strips embedded in an inorganic geopolymeric matrix. the unidirectional reinforcing strip is made of stainless-steel cords, particularly resistant to corrosion, suitable for interventions on substrates subject to rising damp and/or exposure to aggressive environments. the matrix is a polymers-based inorganic mineral with the addition of synthetic fibres, suitable for structural repairs of deteriorated cover concrete being able to be applied with thicknesses between 2 and 40 mm. the geopolymeric matrix also act as inhibitor of corrosion and represents an eco-friendly material for green technology. table 4 presents the mechanical properties of stainless-steel strips and matrix provided in the technical data sheets of the manufacturer/supplier [11, 12]. 3.3. cyclic loading procedure specimens subjected to cyclic loading were tested according to 12 monotonous cycles, as shown in figure 3. the periodic amplitude was selected on the basis of the theoretical yielding loads (fy) of the group c, using four levels of load corresponding around to 25% (fy /4), 33.33% (fy /3), 50% (fy /2) and 75% (3fy /4) of them. from the theoretical calculations, the yield loads of c-a-irs and c-b-irs beams are almost equal. with reference to the theoretical average yielding load of the two strengthened beams and taking into account the applied load before test (self-weight of the beams and the heavy equipment) the values of the load cycles according to the percentages above described were defined. the values of the load cycles correspond to 12, 18, 30, and 42 kn, respectively. the load history was repeated three times for each step and later the beams were monotonically loaded until failure. 3.4. test set-up each beam was subjected to two equal loads applied symmetrically to the mid-span as shown in figure 2. the shear span (a) was 900 mm and 1500 mm for group a and b, respectively. the distance from support to point load was chosen to have the a/d ratio equal to 4.1 in all tested specimens. the load was monotonically applied to the beams of group m by means of a hydraulic actuator and recorded manually using a 300 kn load cell. the specimens of group c were loaded by using hydraulic unit that controls a 140 kn jack and 160 kn load cell, with a frequency acquisition of 2 hz. this equipment allows to apply a load and keep it constant. the load equipment was connected to a rigid steel frame for the application of the load on the elements. the vertical displacements were recorded using three linear variable displacement transducers (lvdts). two were placed at mid-span (both in the front and back side) and one at quarter span. finally, the strain values of all materials (concrete, internal tensile steel reinforcement and external strengthening) were measured by means electrical strain gauges 35 francesco bencardino, mattia nisticò acta polytechnica ctu proceedings figure 3. loading history for the c group beams. � � d�� e�� f� )ljxuh����/rdg�glvsodfhphqw�uhodwlrqvklsv�� figure 4. load-displacement relationships. 36 vol. 33/2022 stainless steel-frcm system for rc beams specimens failure mode fu [kn] εc [‰] εs [‰] εf [‰] ∆fu µδ m-a-cb concrete crushing 37.22 3.67 13.38 8.64 m-a-eb end debonding 43.22 1.50 3.75 1.87 1.16 2.16 m-a-irs intermediate debonding 47.39 2.63 9.81 11.62 1.27 3.73 m-b-cb concrete crushing 76.33 3.68 17.54 4.38 m-b-eb end debonding 89.52 3.53 12.03 5.99 1.17 1.83 m-b-irs concrete crushing 98.96 4.65 15.49 10.73 1.30 3.05 c-a-irs concrete crushing 83.60 1.89 15.31 10.43 c-b-irs intermediate debonding 90.00 1.25 6.24 7.26 table 5. experimental results. � � d�� e� f�� � � g�� h� i�� )ljxuh����([shulphqwdo�idloxuh�prghv�ri�wkh�ehdpv��d��0�$�(%��e��0�$�,56��f��0�%�(%��g��0�%� figure 5. experimental failure modes of the beams: a) m-a-eb, b) m-a-irs, c) m-b-eb, d) m-b-irs, e) c-a-irs and f) c-b-irs. in the middle section of the beams. in the specimens monotonically tested the load was periodically paused in order to identify and mark crack formations and growth while in the cyclic tests it was performed at the end of each step. 4. experimental results and discussion static and cyclic flexural response at failure was experimentally studied. the ultimate load and overall behaviour of all tested beams are analysed and discussed through the comparison of load-deflection curves to measure the influence of the two strengthening methods (eb and irs) on the flexural capacity. the load-deflection curves are shown, for all tested beams, in figures 4a, 4b and 4c. the deflection reported represents the average value measured by lvdts located in the centre of beam span. the classical vertical flexural cracks were observed during the loading and at failure. table 5 summarises the failure modes of all specimens, the maximum load (fu), the strain measured on the top surface of the compressed concrete (εc), on the bottom tension steel bars (εs) and on the stainless-steel strips (εf ). ∆fu indicates the ratio between the ultimate load of strengthened beam and that of control beam; furthermore, for the beams monotonically tested, an evaluation of the ductility in terms of displacement is given as ratio between the deflection in the mid-span section at failure and at yielding of the internal tensile steel (µδ ). from a first analysis, both strengthening systems of group m show an increase of the failure load and a decrease of the ultimate deflection and ductility compared to the control beams (m-a-cb and m-bcb). furthermore, the structural behaviour of the beams m-a-irs and m-b-irs was better compared to the corresponding rc beams strengthened with 37 francesco bencardino, mattia nisticò acta polytechnica ctu proceedings traditional eb technique (m-a-eb, m-b-eb). from the analysis of deformations and failure modes the irs solution provided a greater utilization of all the materials, compared to the traditional eb technique. the failure modes of the strengthened beams with s-frcm strips are shown in figures 5a-f. on the other hand, the beams of group c (c-airs and c-b-irs) shown an excellent behaviour under cyclic loads without premature debonding of the reinforcement system until failure. these first results of group c is a part of a wider experimental programme aimed to investigate the performance of fullscale beams strengthened with the traditional (eb) and innovative (irs) methods in structures affected to variable loads during the service life (such as deteriorated rc deck bridges). 5. conclusions this experimental investigation demonstrated the effectiveness of the irs technique in strengthening of flexural rc members and in reduction of debonding process, showing also an excellent behaviour both under cyclic and monotonic loads. the test results also exhibited that the innovative irs technique had superior flexural performance to the traditional eb strengthened technique, particularly when considering the economic, eco-friendly, and of environmental sustainability viewpoint. in detail, the results showed that with eb and irs technique, the average flexural strength of the beams is improved by around 15% and 30% respectively, compared to control beam under static loading; and great behaviour under cyclic loading for the beams strengthened by means the irs 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[10] c. menna, d. asprone, c. ferone, et al. use of geopolymers for composite external reinforcement of rc members. composites part b: engineering 45(1):1667-76, 2013. https: //doi.org/10.1016/j.compositesb.2012.09.019. [11] kimisteel inox st5-0218, 2019. https://www.kimia.it/sites/default/files/docs/ st/it-kimisteel-inox.pdf. [12] betonfix monolite n st6-0319, 2019. https://www.kimia.it/sites/default/files/docs/ st/it-betonfix-monolite-n.pdf. 38 https://doi.org/10.1016/j.conbuildmat.2016.05.053 https://doi.org/10.1061/(asce)cc.1943-5614.0000960 https://doi.org/10.1016/j.compositesb.2016.03.022 https://doi.org/10.1061/(asce)cc.1943-5614.0000218 https://doi.org/10.1061/(asce)cc.1943-5614.0000882 https://doi.org/10.1016/j.compositesb.2018.04.051 https://doi.org/10.3390/fib8020008 https://doi.org/10.1016/j.engstruct.2019.109348 https://doi.org/10.1016/j.istruc.2019.02.005 https://doi.org/10.1016/j.compositesb.2012.09.019 https://www.kimia.it/sites/default/files/docs/st/it-kimisteel-inox.pdf https://www.kimia.it/sites/default/files/docs/st/it-betonfix-monolite-n.pdf https://doi.org/10.14311/app.2022.33.0175 acta polytechnica ctu proceedings 33:175–180, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague bending behavior of adhesively-bonded engineered wood-concrete composite decks qiuni fua, libo yanb, ∗, bohumil kasalb a technische universität braunschweig, institute of building materials, concrete construction and fire safety, department of organic and wood-based construction materials, hopfengarten 20, 38102 braunschweig, germany b fraunhofer insitute for wood research wilhelm-klauditz-institut wki, bienroder weg 54e, braunschweig 38108, germany ∗ corresponding author: l.yan@tu-braunschweig.de, libo.yan@wki.fraunhofer.de abstract. four-point bending tests were conducted on five medium-sized (i.e., 2300 mm in length and 215 mm in width) engineered timber (laminated veneer lumber (lvl) and cross-laminated timber (clt)) concrete (wood chip concrete and plain concrete) composite decks. the concrete was glued to the wood substrate with epoxy and polyurethane adhesives. the observed failure modes of the composite decks were concrete crushing or wood failure in tension or shear. no failure of the adhesive interface was observed and the decks behaved linearly until failure. in the subsequent analysis, the authors quantified the shear flexibility of transverse layers (stressed perpendicular to the fiber direction) in clt and lvl boards and its effect on the bending stiffness on the composite decks using γ-method described in en 1995-1-1 (ec5). the analytical predictions of the effective bending stiffness were verified via experiments, showing consistently good agreement. keywords: γ-method, adhesive bond, bending behaviour, timber-concrete composite, wood chip concrete. 1. introduction timber-concrete composite (tcc) structural elements have advantages of higher stiffness, fire resistance and better acoustic performance compared with massive timber. connection between concrete and wood is usually achieved by conventional mechanical dowel-type fasteners (shear studs) or notches. mechanical fasteners are not rigid and there is always a slip between the wood and concrete. thus, full composite action of tcc connected by mechanical fasteners is hardly achievable. adhesive bonding is often considered as rigid. the application of glued wood-concrete connections is not fully explored and the knowledge in this area is limited [1–7], let alone the relevant specifications in standards. studies [1–7] demonstrated that glued tcc systems had higher stiffness (e.g. slip modulus of connections and bending stiffness of tcc beams) compared to mechanically connected ones, and sufficient adhesion strength for structural applications under short-term loading and ambient-temperature. the wood products involved in these studies [1–7] were solid timber, glulam or vertically used laminated veneer lumber (lvl) assuming the rigid characteristics of the used adhesives, the behavior of these kinds of tcc beams was characterized by assuming continuous strain at the wood-concrete interface and using the known transformed cross-section procedure [5, 8– 10].the low shear modulus of transverse layers in the clt and lvl boards subjected to bending can cause significant shear deformation (or slip) and hence increase the deflection. in the present work, the authors studied the shear behavior of glued cltand lvl-concrete connections [11]. they further experimentally investigated the bending behavior of cltand lvl-concrete composite decks. specifically, clt and lvl boards were bonded wood chip concrete (wcc) and with plain concrete (pc) using epoxy and polyurethane (pur) and adhesives. the addition of wcc [12] resulted in reduced concrete density and increased thermal insulation properties of plain concrete. the polyurethane (pur) used in the present study was stiff, its modulus of elasticity was as high as that of the used epoxy (around 3.5 gpa), showing nearly no deformability in the adhesive layer. the pur in this study was different from the pur in ref. [13] (modulus of elasticity: 0.008 gpa) that resulted in visible slip at bond line (over 1 mm). to describe the slip in transverse layers and its effect on the bending stiffness of composite decks, the authors used the shear analogy method [14] for quantifying shear deformation of the transverse layers in combination with γ-method of ec 5 [15]. the predictions of the effective bending stiffness were compared with the experimental results, showing a good match. 175 https://doi.org/10.14311/app.2022.33.0175 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en q. fu, l. yan, b. kasal acta polytechnica ctu proceedings clt flatwise lvl thickness (mm) 80 (20 + 40 + 20) 69 (thickness of each veneer: 3 mm) shear modulus (mpa) 50 (rolling shear, gr ) 22 (perpendicular to grain, g90) elastic modulus in grain direction (e0) (gpa) 12 11.4 table 1. mechanical properties of the engineered wood products [16]. 2. experimental studies 2.1. overview of test plan engineered timber, clt and lvl, and concrete, wcc and pc used in this study were identical to those in the previous study [11]. epoxy and pur were selected due to their bonding performance [11]. the authors constructed five mediumsized specimens, wcc-epoxy-lvl, wcc-pur-lvl, wcc-epoxy-clt, wcc-pur-clt and pc-epoxylvl, and tested them under bending. figure 1 (e) shows the configurations of the specimens. 2.2. materials the mechanical data and configurations of clt and lvl are shown in table1. moisture content of the used wood was about 10%. the beech wood chips used in wcc were in a relatively flat shape, having around 20 mm side length and 2.4 mm thickness. the density of the wood chip was 0.68g/cm3. the mass ratio of the constituents of wcc was cem ii/b-s 42.5 r cement: water: 0-2 mm sand: 2-8 mm gravel: 8-16 mm gravel: wood chip = 1: 0.60: 2.00: 1.27: 2.32: 0.11. the mass ratio of the constituents of pc was cem ii/b-s 42.5 r cement: water: 0-2mm sand: 2-8 mm gravel: 8-16 mm gravel = 1: 0.60: 2.00: 1.27: 2.72. the compressive strength of the wcc and pc cylinders were 25.8 mpa (sd = 3.79 mpa) and 27.9 mpa (sd = 1.73 mpa) based on five samples, respectively. the respective elastic modulus of pc and wcc were 29.9 gpa and 12.3 gpa. according to the product of primetm 20lv from [17] datasheet, the elastic modulus of epoxy is 3.5 gpa. the pull-off tests in authors previous study [11] showed that the load-displacement curves of the used epoxy and pur almost overlapped at the initial loading stage, so the elastic modulus of the used pur is similar to that of the epoxy. 2.3. specimen preparation when the concrete was cured for 28 days, the concrete surfaces were cleaned and glued to shaped wood decks with wood decks on the top for the first 3 days. the specimens were turned upside down since the fourth day and tested after the seventh day. the detailed processes can be found in ref. [16]. 2.4. experimental setup a load (f ) was quasi-statically applied on the tcc decks with a rate of 0.35 mm/s through a two-point spreader, as shown in figure 2. it also shows the locations of the support and loading points. a deflection at the mid span was measured by two lvdts (linear variable differential transducers). 3. results and discussion 3.1. load-carrying capacities the load versus mid-span deflection curves of all the tcc decks showed linear behaviour until the ultimate failure. figure 3 shows the response of wccpur-clt as an example, and the ultimate loads, the ultimate mid-span deflections and bending stiffness (ei)test of all the specimens are summarized in table 3. the calculation of (ei)test was based on the ascending slope of a load versus mid-span deflection curve as follows [16]: (ei)ef = f w a ! 3l2 − 4a2 " 48 (1) where w is the average reading value of the two lvtds; a is the shear span; and l indicates the span of the deck. from table 3, it follows that the bending stiffness of a beam was not affected by adhesive type. the reason was that a rigid connection was achieved by using either epoxy or pur, which was also demonstrated by the absence of slippage between concrete and wood decks (see section 3.2 and table 3). the specimens comprising clt boards were stiffer than the counterparts made of lvl (table 3). the inherent reasons are discussed in section 4. on the contrary, the ultimate load-bearing capacities of the lvl-specimens were higher than that of clt-specimens., due to different failure modes (table 3). as the pc had higher elastic modulus and strength compared with wcc, pc-epoxy-lvl had the largest flexural stiffness (4.12 × 1011n.mm2) and ultimate load-carrying capacity (63.2 kn) [16]. the average bending stiffness and load-bearing capacity of wcc-epoxy-lvl and wcc-pur-lvl was around 78.9% and 70.0% of those of pc-epoxy-lvl, respectively. meanwhile, the thermal conductivity of wcc could be only 68.9% of that of pc. [18]. besides, wide application of wcc can reuse wood waste, and reduce the demand for natural aggregates. the load-bearing capacity of the weakest specimen, wccepoxy-clt, was 39.1 kn, equivalent to a uniformlydistributed load of 85.4 kn/m2. considering serviceability limit design, ec 5 [15] allows the maximum 176 vol. 33/2022 bending of woodconcrete composite deflection of a slab to be l/300 (i.e. 7.1 mm). this deflection corresponds to an equivalent slab load of 34.7 kn/m2 in the case of wcc-epoxy-clt. 3.2. failure modes the failure modes of all the specimens are summarized in table 3. neither of the specimens was failed by adhesive bonding interface, indicating that the used adhesives can provide satisfactory bonding performance. the different mechanical properties of clt and lvl caused different failure modes of the corresponding tcc decks. -see [16] for details. one can also have a deeper appreciation of the structural behaviour of lvl and clt in tcc decks in section 4. 4. prediction of bending stiffness in building slab design, deflection is one of the important serviceability considerations. structural engineers usually calculate deflections with eulerbernoulli beam theory that does not take shear deformation into account. however, when engineered wood (e.g. clt) is under bending, transversely stressed layers have significant shear deformation, because of low rolling shear modulus (gr ) and low shear modulus in the perpendicular direction (g90) (table 1). to consider the shear flexibility of transverse layers, the authors regarded it as sources of partial composite action and calculated the effective bending stiffness (ei)ef for all the specimens based on γmethod in ec 5 [15]. the γ-method [15] allows a composite cross-section to be comprised of three parts with two connecting interfaces featured with the respective degree of composite action (γi) relative to the middle layer (referred as "part 2" in figure 4 (a)). a term γi ranges between 0 (no interaction) and 1 (rigid connection). then (ei)ef of a composite deck is calculated with equation (2) [15]. (ei)ef = 3# i=1 ! ei ii + γi ei ai a2i " (2) where ei is the modulus of elasticity of part i in the longitudinal direction; the term ii means moment of inertia of part i; the term ai denotes the cross-section area of part i; the term ai indicates the distance between the neutral axis (i.e. n.a. in figure 4 (a)) and the centroid of part i. according to ec5 [15]: γi = $ 1 + π2 ei ai si % (ki l2) & −1 (i = 1 and 3) (3) where ki is the slope of a shear force slip curve (i.e. slip modulus) of a connection, and si denotes the equidistance between shear fasteners connecting part 2 and part i. for a continuous adhesive connection, ki can be defined based on a given length, e.g. 1 mm. then, si is equal to 1 mm. as explained in ref. [16], the clt board (part 2) and concrete (part 1) were rigidly bonded based on the experimental observations (i.e. k1 = ∞ γ1 = 1). part 2 is the reference layer, so γ2 is equal to 1. in this paper, the slip modulus k3 relates to the shear connection between the top (part 2) and bottom (part 3) longitudinal layers of clt. therefore [16], k3 = gr · b · (1000 mm) / htr (4) where gr denotes the rolling shear modulus of the clt ( refer to table 1), htr denotes thickness of the transversal layer, and b is the width of the decks. according to ec [15]: a2 = γ1 e1 a1(h1 + h2) − γ3 e3 a3(h2 + h3) 2 3' i=1 γi ei ai (5) the terms a1 and a3 can be obtained based on the geometric relationship in figure 4 (a). to model the shear flexibility of dispersed transversal veneers in lvl, the longitudinal and transversal veneers were gathered into the respective two layers for simplification. the full concept is illustrated in figure 4 (b). for details regarding the modelling and analysis, please refer to the authors’ previous work in ref. [16]. in simple terms, equations (2) through (5) are applicable for tcc decks using lvl, with h1 = 0 and g90 replacing gr . all the detailed derivations can be found in ref. [16]. for comparison, γi and (ei)ef of tcc decks comprising the respective clt and lvl boards are presented in table 2. the degree of composite action in the tcc decks comprising lvl (0.52) was lower compared with that of the tcc decks comprising clt (0.70). the reason was that the transversely stressed veneers have a very low shear modulus (22 mpa). the (ei)ef of cltinvolved tcc decks was slightly higher than that of lvl-involved tcc decks, which agrees with the test results (table 3). the higher (ei)ef of clt-involved tcc decks stemmed from a larger γ3 and a larger a3 (see figure 4). the specimen made of pc had the largest stiffness due to the larger elastic modulus of the plain concrete. the larger elastic modulus of pc led to the largest stiffness of pc-epoxy-lvl. as shown in table 2, the (ei)ef calculated in this study conservatively predicted the corresponding test results with an error of around 10% [16]. ignoring the shear deformation (i.e. γ3 = 1), ei was overestimated by around 34% on average. with the latter kind of calculation, another undesired situation is that the errors of predictions of ei varied significantly (cov = 14.1%), depending on different types of engineered wood products having different degrees of shear flexibility. this problem was solved by this study through quantifying the shear flexibility of transversal wood layers [16]. the errors of predicted (ei)ef of cltand lvl-involved tcc decks had small variations 177 q. fu, l. yan, b. kasal acta polytechnica ctu proceedings figure 1. dimensions of fabricated tcc decks. concrete lvdtȭ1ȱ(2) 710ȱmm 710ȱmm 710ȱmm timber f/2 f/2 f355ȱmm 355ȱmm roller figure 2. setup of 4-point bending test. specimen id γ1 γ3 ei a (n.mm2) ei/(ei)test (ei)ef (n.mm2) (ei)ef /(ei)test wc-epoxy-clt 1 0.70 4.13×1011 1.14 3.21×1011 0.89 wc-pur-clt 1 0.70 4.24×1011 1.13 3.33×1011 0.89 wc-epoxy-lvl \ 0.52 4.62×1011 1.42 2.90×1011 0.89 wc-pur-lvl \ 0.52 4.90×1011 1.50 3.05×1011 0.93 pc-epoxy-lvl \ 0.52 6.15×1011 1.49 3.82×1011 0.93 average 1.34 0.91 coefficient of variance 14.0 % 2.6 % a the term ei denotes the flexural stiffness of tcc decks, ignoring the shear flexibility of transvers layers. table 2. predictions of flexural stiffness. figure 3. load vs. displacement responses of wccpur-clt at mid-span. (cov = 2.6%) (table 2). thus, this study provides an effective solution for the calculation of (ei)ef (or deflection) of tcc decks using engineered wood. 5. conclusions from this study, some conclusions can be drawn, as follows: 1. the purand epoxy-bonded wood and concrete could form a strong and rigid interfaces to transfer the shear in tcc decks subjected to quasi-static bending, without failure and visible slip. 2. the failures of tcc decks in this study were brittle, which were governed by wood fracture and concrete crushing. 3. wood chip concrete is feasible to be used with the engineered wood as composite structural slabs. engineered wood could be used together with wcc to realize a sustainable solution, with satisfactory structural performance. 4. the method in this study provides an effective solution for the calculation of (ei)ef (or deflection) of tcc decks under bending. acknowledgements the research was financially supported by fachagentur nachwachsende rohstoffe e. v. (fnr, agency for renewable resources) founded by bundesministerium für ernährung und landwirtschaft (bmel), under the grant award no.: 22011617, and by bundesministerium für bildung und forschung (bmbf) (grant no.: 031b0914a). 178 vol. 33/2022 bending of woodconcrete composite (a) concrete-clt composite decks concreteȱ(ec) adhesive* 40 20 1000 ȱlongitudinalȱlayer ȱlongitudinalȱlayer transversȱlayer gr 20 e0 e0 3ȱ (5 ) 50 h t r h 2 h 3 h 1 model n.a. a 1 a3 h2 /2 h1 /2 h3 /2 a 2 *ȱtheȱmomentȱofȱinertiaȱofȱtheȱadhesiveȱlayerȱisȱignored,ȱbutȱitsȱthicknessȱisȱconsideredȱwhenȱdeterminingȱgeometricȱpositionsȱforȱotherȱmaterialȱlayers. construction concreteȱ(e1) adhesive 40 20 1000 ȱlongitudinalȱlayer ȱlongitudinalȱlayer k3=gr*b*(1000mm)/htr partȱ1 partȱ2 partȱ3 20 k1=infinite e3 e2 3ȱ (5 ) 50 (b) concrete-lvl composite decks partȱ3 k3=g90*b*(1000mm)/htr h 3 =5 4 h 2 3ȱ (5 ) ȱrigidȱadhesiveȱlayeradhesive:ȱrigidȱbond htr=15ȱmm model construction concreteȱ(ec) ȱlvlȱ(e0,ȱg90) ȱlongitudinalȱlayertransversȱlayer 3 3 ȱlegends 69 1000 3ȱ (5 ) 50 concreteȱ(e2) ȱlvlȱ(e3) 34 ,5 1000 50 34 ,5 27 27 centroidȱline ofȱwood partȱ2 figure 4. analytical model of composite decks (unit: mm) quoted from ref. 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[8] dias a m p g, mechanical behaviour of timber-concrete joints. 2005, technische universiteit delft: netherland. [9] l. boccadoro, r. steiger, s. zweidler, et al. analysis of shear transfer and gap opening in timber-concrete composite members with notched connections. materials and structures 50(5), 2017. https://doi.org/10.1617/s11527-017-1098-3. [10] v. bajzecerová, j. kanócz. timber-concrete composite elements with various composite connections. part 2: grooved connection, wood research, 59(3):27-638, 2014. 179 https://doi.org/10.1617/s11527-006-9154-4 https://doi.org/10.4028/www.scientific.net/amr.969.155 https://doi.org/10.18848/2154-8587/cgp/v03i03/37394 https://doi.org/10.1061/(asce)st.1943-541x.0000228 https://doi.org/10.1061/(asce)st.1943-541x.0000251 https://doi.org/10.1002/mawe.201300026 https://doi.org/10.1002/bate.201110017 https://doi.org/10.1617/s11527-017-1098-3 q. fu, l. yan, b. kasal acta polytechnica ctu proceedings [11] q. fu, l. yan, t. ning, et al. interfacial bond behavior between wood chip concrete and engineered timber glued by various adhesives. construction and building materials 238, 2020. https: //doi.org/10.1016/j.conbuildmat.2019.117743. [12] t. ning. interfacial bond and structural performance of adhesively-bonded light-weight timber-wooden chip concrete composite structures, in institut für baustoffe, massivbau und brandschutz (ibmb), technische universität braunschweig: braunschweig, germany, 2018. [13] j. r. cruz, s. serȩga, j. sena-cruz, et al. flexural behaviour of nsm cfrp laminate strip systems in concrete using stiff and flexible adhesives. composites part b: engineering 195, 2020. https: //doi.org/10.1016/j.compositesb.2020.108042. [14] q. fu, l. yan, t. ning, et al. behavior of adhesively bonded engineered wood wood chip concrete composite decks: experimental and analytical studies. construction and building materials 247, 2020. https: //doi.org/10.1016/j.conbuildmat.2020.118578. [15] p. mestek, h. kreuzinger, s. winter. design of cross laminated timber (clt), in 10th world conference on timber engineering, miyazaki, japan, 2008. [16] gurit. primetm 20lv-epoxy infusion system, 2018. https://www.gurit.com/sitecore/content/old-pro duct-pages/other-products/laminating-infusio n-systems/prime-20lv. [17] en 1995-1-1, eurocode 5 design of timber structures part 1-1: general common rules and rules for buildings. european committee for standardisation, brussels, 2004. [18] m. bederina, l. marmoret, k. mezreb, et al. effect of the addition of wood shavings on thermal conductivity of sand concretes: experimental study and modelling. construction and building materials 21(3):662-8, 2007. https: //doi.org/10.1016/j.conbuildmat.2005.12.008. 180 https://doi.org/10.1016/j.conbuildmat.2019.117743 https://doi.org/10.1016/j.compositesb.2020.108042 https://doi.org/10.1016/j.conbuildmat.2020.118578 https://www.gurit.com/sitecore/content/old-product-pages/other-products/laminating-infusion-systems/prime-20lv https://doi.org/10.1016/j.conbuildmat.2005.12.008 https://doi.org/10.14311/app.2022.33.0146 acta polytechnica ctu proceedings 33:146–152, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague advanced hybrid nonmetallic composite reinforcement for concrete structures vyacheslav ruvimovich falikmana, ∗, dmitry anatolievich ilyinb, valentina fedorovna stepanovaa a scientific research center "construction"/scientific research institute for concrete and reinforced concrete, department on structural concrete durability, 2nd institutskaya st., 6, 109428, moscow, russia b moscow state university of civil engineering (national research university)/ institute of construction and architecture, department of building materials, yaroslavskoye shosse, 26, 129337, moscow, russia ∗ corresponding author: vfalikman@yandex.ru abstract. during the last decades, fiber reinforced polymer (frp) reinforcing bars for concrete structure has been extensively investigated and a number of frp bars became commercially available. however, major shortcomings of the existing frp bars are its low elastic modulus and high initial cost compared to conventional steel bars. the possibility to obtain a hybrid composite reinforcement (hcr) with increased performance based on glass and carbon fibers (gcfrp) is considered. the optimal content of carbon fibers in the amount of 6.3 − 6.5 % of the mass of the hcr was established. further increase in the carbon fiber content gives a slight improvement in physical and technical characteristics, which is not comparable to the increase in the cost of the material. the manufacturing technology of hcr has been developed. the effect of hybridization on tensile properties of frp bars were obtained by comparing the results of tensile test with those of non-hybrid gfrp bars. operation regularities of hcr in the bent concrete beams are established. hcr can increase the stiffness of concrete beams by 15 % and crack resistance by 12 % in comparison with glass composite reinforcement. dependences for predicting the hcr elasticity modulus are established. physical and technical characteristics of hcr, including adhesion to concrete and resistance to the alkaline medium, were established. high durability of hcr for more than 50 years is experimentally shown. experimental-industrial concrete piles, reinforced with gcfrp bars were produced and tested. for further development, new types of hcr, as well as a study of prestressed concrete structures are recommended. keywords: durability, elasticity modulus, fiber reinforced polymer (frp), glass and carbon fibers, hybrid composite reinforcement (hcr). 1. introduction the russian industry for the production of composite polymer reinforcement (frpr) is rapidly developing. non-metallic reinforcements are becoming valuable alternatives for the classical concrete reinforcements [1]. this evolution started in the late eighties of last century. having high tensile strength, low density, dielectric properties and high chemical resistance, frpr is of great interest to builders and researchers. composite materials of different type were introduced to overcome the inherent disadvantages of concrete: strength, durability, corrosion of steel rods, creep and shrinkage, flexibility and retrofitting aspects. glass, aramid and carbon are the most popular fiber reinforcements, embedded in a pure polymer or in a polymer modified matrix. material properties, reinforcement design, analysis and layout became important research topics, but practical applications came immediately after research, and sometimes even preceded research. along with the positive properties, frpr has some disadvantages. among them, it’s necessary to mention a low modulus of elasticity and low heat resistance, increased creep under prolonged static load. all this hinders the widespread application of composite rebar in the construction practice. some of the above disadvantages could be eliminated when modifying the epoxy binder with nanoadditives. in the last decade, nanoparticles of various natures have been tested for modification of polymeric materials, wiz carbon nanotubes, fullerenes, astralenes, ultrafine diamond powders, nanodiamonds, montmorillonite, aerosil, schungite, metal-containing nanocomposites, etc. [2–4]. the positive effect of the modification by copper and nickel nanoparticles was confirmed by one of the russian frpr manufacturers. they achieved stable increased physical and mechanical properties of the reinforcement and its high corrosion resistance [5]. however, the modification of the polymer binder is not enough to significantly increase the elastic modulus of the frpr. this is due to the initial low modulus of elasticity of the glass fiber (70-80 gpa only). in turn, carbon fiber (cf) has a high modulus of elasticity (220 gpa), but the cfrp reinforcement is of 25-30 times more expensive. the optimal solution may be a hybrid com146 https://doi.org/10.14311/app.2022.33.0146 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 hybrid nonmetallic composite reinforcement � �� �� figure 1. experimental samples of frpr; cf location. seq no. component kind quantity [%] 1 fibreglass jsc "fiberglass" 71.3 2 carbon fiber aksa a−49 3.7 3 multicomponent epoxy binder solidifier−imthfa1 25 promoter−tdmamph2 25 1 isomethyl tetrahydrophthalic anhydride 2 2, 4, 6−tris−(dimethylaminomethyl)−phenol table 1. composition of hcr experimental samples designation nominal cross density tensile ultimate elasticity unit bond diameter area strength modulus extension resistance dn ρ σb ef εb [mm] [mm2] [kg/m3] [mpa] [mpa] [%] [mpa] gfrp 11.89 111.03 2025 935 51500 1.72 11.92 acc−c 11.98 112.72 2009 1115 59500 1.87 13.90 acc−k 11.89 111.03 2046 1110 62000 1.60 12.76 table 2. test results posite reinforcement (hcr) based on carbon and glass fibers, impregnated with a nanomodified epoxy binder. the use of two fiber types allows increasing of the elasticity modulus. nanomodification of the epoxy binder provides high-quality joint work of the "fiber + binder" system. the result of this processing is a dense structure of the reinforcement, an inseparable relief from the bar, and a high proportion of reinforcing fiber. 2. experimental and discussion 2.1. making of hcr and its basic material properties previously, the hybrid effect was studied on the example of composite reinforcement made of glass and carbon fibers with their different locations. the authors noted that the final strain under tension of bars consisting of 37 % glass fibers and 23 % carbon fibers increased by 3 − 33 % compared to the carbon frpr [6]. in the course of experimental work, a hybrid composite rebar (hcr) was manufactured and tested, in which 3.73 − 6.3 % of the glass fibers were replaced by cf with a modulus of elasticity of 220 gpa. at the same time, experimental and conventional samples of gfrp bars with a diameter of 12 mm were studied. samples of hybrid composite (hcr) and glass composite rebar (gfrp bar) were produced using industrial line. in experimental samples, 3.7 % of the glass fibers were replaced with cf. the theoretical value of the elasticity modulus of hybrid reinforcement calculated in accordance with voigt’s formula is 63.9 gpa. at the same time, two schemes were adopted for the location of the cf (figure 1): 1−cf was concentrated in the center of the composite; the designation of this type of reinforcement is acc−c; 2−cf were evenly distributed along the profile outline; the designation in this case is acc−k. table 1 shows the ratio of hybrid composite reinforcement components. as the primary roving, optical fiber of jsc "fiberglass" located in gus khrustalny was selected. carbon fiber used was aksa a-49 brand. manufacture of frpr was carried out according to the needletrusion technology [7]. the main performance characteristics of experimental samples according to gost 31938-2012 are given in table 2. the rupture fracture pattern of the frpr sample is shown in figure 2a. the dependence of the strain on the stress (load) during the tensile test is shown in figure 2b; it shows repeated peak stresses, which 147 v. r. falikman, d. a. ilyin, v. f. stepanova acta polytechnica ctu proceedings are explained by the hybrid effect. the dependence of deformations on stress (load) during tests for axial expansion of acc−k−10 samples with a carbon fiber content of 6.3 % is also characterized by repeated peak stresses, which indicates the pseudo-plastic behavior of combined composite reinforcement samples at break. the maximum strain of 1311 mpa is reached at 1.93 % deformation, after which the strain drops to 1200 mpa with a further rise to 1250 mpa. then there is a breakage of the test sample acc−k. the deformation reduction is occurred normally in accordance with the "mixtures law". the substitution of 6.3 % of glass fibers with carbon fibers resulted in an strength increase of the reinforcement by 17.4 % and the modulus of elasticity by 31.5 % compared to the samples of gfpr bars. the quality of frpr depends, first of all, on its components used in the production, viz fiber, multicomponent polymer binder, and their ratio in the matrix of the composite material. in addition, a significant role is played by production technology of frpr, namely the tension force of the fiber, temperatures of polymerization of the binder, and cooling the finished product etc. selecting of the optimal frpr composition and the observance of technological processes allows obtaining a product with the specified parameters. 3. corrosion (deterioration) resistance of hcr to alkaline environment when achieving high physical and mechanical properties of frpr, the determining factor becomes its corrosion (deterioration) resistance to aggressive media. in russian [8–12] and foreign [13–15] studies, the aggressive medium for composite polymer materials is usually taken as an alkaline medium, which is being characteristic for wet concrete. the aggressiveness of this medium is estimated by the ph value in the range of 12-14. to simulate the liquid environment of concrete, wet concrete may used directly [11]; aqueous naoh solutions of various concentrations [10], saturated water solutions ca(oh)2 with ph=13 [15] or ph=12 [14] may be used also. in order to accelerate the aggressive action of the alkaline medium on the material under study, the medium is heated to 50◦c [14], 55◦c [11], 70◦c, 80◦c [13], 100◦c [8]. to date, however, there are no generally recognized theoretical dependencies that allow predicting the durability of the composite material on the basis of accelerated tests. thus, according to the authors [13], 50 years of operation may be imitated by holding of the bar for 28 days in a saturated ca(oh)2 solution with ph=13 at a temperature of 80◦c. in [14], samples of frpr were kept in a container with an alkaline medium at a temperature of 50◦c for 52 weeks. the authors believe that 1 day under such conditions corresponds to 101 days under natural operating conditions. test results show that the service life of the frpr in concrete is about 100 years. in [15], tests were performed on frpr samples placed in a saturated solution of ca(oh)2 with ph=12 at a temperature of 50◦c. calculations made by authors show that 1 day of such exposure is comparable in intensity to more than 100−day alkaline exposure in concrete. thus, the tested gfrp bars are able to serve in concrete for more than 47 years. in this work, the assessment of the hybrid composite reinforcement (hcr) stability to the alkaline environment was carried out according to gost 31938. preliminary tests of the reinforcement for axial tension showed that the samples of hcr with the edge arrangement of carbon fibers, akk-k series (figure 1), more efficient than with the central placement, taking in account of elasticity modulus increasing. when replacing 3.7 % of glass fibers with carbon fibers, the modulus of elasticity was increased to 18 %. in this regard, samples of gfrp bars and acc-k were studied. an alkaline solution with a ph of 12.6−13 was used as an aggressive medium in the study, since this medium is the closest to the composition of the liquid concrete phase. the alkaline solution had the following composition: 8.0 g naoh and 22.4 g koh per 1 liter of distilled water. the samples were immersed in an alkaline solution and kept in a closed container under a constant solution temperature of 60 ± 3 ◦c for 30 days, with only the middle section of the samples exposed to the alkaline environment; during the tests, the appearance of the samples was monitored before and after aging in an alkaline solution (figure 3 a, b). the ph value of the solution was maintained in the range of 12.6−13.0 during the entire experiment. after exposure, tests were performed to determine the axial tensile strength in accordance with gost 32492−2013. the results are presented in table 3. after aging in an alkaline environment, samples of frpr are observed: rupture of the profile-forming thread, fading, and the appearance of «whitishness» (figure 3). for gfrp samples, the rupture resistance decrease was by 5.2 %, for the akk-k samples it was 5.6 %. the reduction of the elastic modulus was obtained by 2.31 % for gfrp samples and 1.26 % for acc−c samples. interstate standard gost 31938-2012 did not specify an elastic modulus reduction of the frpr after exposure to an alkaline environment. so, it seems reasonable to make provision for this limitation during next revision of gost, as it was done in official edition for the strength degradation of the frpr under axial tension. figure 4 shows samples of acc−k tested for rupture after exposure to an alkaline medium. the samples were destroyed in the central part with a characteristic "fluff out" of the fibers. 148 vol. 33/2022 hybrid nonmetallic composite reinforcement ɚ�� e�� figure 2. frpr tests. a) the nature of frpr destruction at the break; b) tensile stress-strain diagram. ultimate tensile stress elasticity modulus, ef seq series reference value after exposure reference value after exposure no. frpr unit average unit average unit average unit average value value value value ∆1 value value value value ∆1 [mpa] [%] [gpa] [%] 1 gfrp 935.5 935.1 861.8 886 5.2 52.7 53.20 50.7 52.0 2.31 2 943.6 844.5 55.1 53.4 3 932.0 887.6 52.6 50.7 4 938.1 879.0 54.7 51.0 5 930.2 991.1 51.3 55.8 6 931.4 853.2 52.8 50.2 7 acc−k 1080.1 1085.1 1040.4 1024 5.6 63.9 62.82 59.7 62.0 1.26 8 1100.2 992.3 62.3 61.3 9 1050.4 1074.2 64.1 58.9 10 1090.6 986.7 61.9 64.6 11 1079.2 998.4 62.9 62.8 12 1110.1 1052.4 61.8 64.9 table 3. frpr strength under axial tension ɚ�� e�� figure 3. appearance of acc−k samples before (a) and after (b) exposure in an alkaline medium 149 v. r. falikman, d. a. ilyin, v. f. stepanova acta polytechnica ctu proceedings figure 4. acc−k samples after the rupture test figure 5. concrete beam reinforcement diagram 3.1. tests of concrete beams reinforced with hcr to study a behavior of the hcr in concrete structures, beams were made with a cross section of 120 × 220 × 1200 mm. prototype beams were made of ordinary concrete (strength class b 20; cement grade cem 1 42.5). as course aggregate, granite crushed stone was used (fraction of 5−25 mm); fine aggregate was quartz sand. tests of control samples (concrete cubes) were carried out in accordance with gost 10180. the beams were tested for short-term action of bending moments in order to determine the maximum breaking moment, as well as to establish the nature of the collapse, the development of cracks, patterns of development of deflections, and deformations of concrete depending on the type of frpr. the reference structures were reinforced with gfrp bars with a diameter of 10 mm (series b−1); experimental ones were reinforced with hcr having a diameter of 10 mm and a carbon fiber content of 6.3 % (series b−2−acc). the reinforcement scheme for testing of both beams series was adopted the same according to figure 5. bent elements made of smooth steel reinforcement of class a240 with a diameter of 6 mm were used as clamps. during the tests, the span between the supports was chosen of 106 cm; the distance from the supports to the point of the load application was 35.5 cm, and the distance between the clamps was accepted of 70 mm. deflections of concrete beams were measured with a digital deflection meter. according to the readings of the test machine sensors, the movements of the traverse were determined. data were recorded after the application of the load at each stage. concrete deformations were recorded using a multichannel measuring system by tokyo sokki kenkyujo co., ltd., japan (tds−530 model). as measuring instruments, conductive strain gauges with a polyester-based pl-60-11 type substrate with a measuring base length of 60 mm were used. sensors were applied by a special technique to a pre-prepared concrete surface. measurements of concrete deformation and deflection of beams were carried out until the load-bearing capacity of the beams was exhausted. the moment of crack formation and the kinetics of their development were controlled using a special template. the widths of normal cracks were fixed at the level of the stretched zone of the reinforcement, and inclined cracks in the middle of the beam section. the length of the cracks was measured after testing by the marks applied to the beams. the testing process was accompanied by photography. after the destruction of the beams, the destruction zone was fixed; the height of the cracks, the distance 150 vol. 33/2022 hybrid nonmetallic composite reinforcement between the cracks, and the thickness of the concrete protective layer were measured. the actual bearing capacity of the beams reinforced with gfrp bars and hcr was the same amounted to 89 kn, which is of 35 % higher than calculated for the ultimate limit state (uls). for both series of beams, the destruction occurred along inclined sections with a rupture of the frpr in the stretched zone. for beams reinforced with gfrp bars, the first cracks formation occurred under a load of 15 kn, which is of 12 % less than the estimated moment of crack formation. in turn, in the beams with hcr, the first cracks were formed at a load of 17 kn, corresponding to the calculated load. the normative crack opening width equal to 0.7 mm was formed in beams reinforced with hcr under a load of 37 % higher than in beams reinforced with gfrp bars. the longitudinal deformation of compression and extension at height of normal beams section reinforced with a gfrp bars and hcr are linearly with the maximum compression strain at the upper edge and the maximum deformations at the lower edge. at the same time, the deformation of the lower stretched zone of the beams with gfrp bars at maximum load is of 70 % higher than the deformation of the beams with hcr. the maximum deflections of beams reinforced with a hcr are of 15 % lower than the deflections of beams reinforced with gfrp reinforcement. tests of gfrp reinforcement for fire resistance according to russian standards have shown that it is highly inflammable, but non-burning. when testing concrete structures reinforced with gfrp bars, even a short-term combustion occurrence does not occur. the structures are fireproof. 4. conclusion the paper deals with important material properties of frps, on which application and durability as reinforcement based. thus, a hybrid composite reinforcement based on glass and carbon fibers was produced, which is resistant to the alkaline concrete environment. when replacing 3.7 % of glass fibers with carbon fibers, the elastic modulus of the frpr is increased to 18 %, and the elastic modulus remains high (62 gpa) even after exposure to an alkaline environment. analysis of the deterioration nature of samples and test results showed that the edge location of the cf is the most effective. a large convergence of the theoretical (calculated) modulus of elasticity and the test results was achieved; the discrepancy limit does not exceed of 3 %. theoretical calculations and tests have shown that the elastic modulus of the frpr increases with an increase in the content of cf and with a decrease in the amount of polymer binder. at the same time, the feasibility of nanomodification of an epoxy binder with varying amounts of cf was proved. the decrease in the strength of the reinforcement in the axial tension of hcr after exposure to an alkaline environment is not key question, since its strength is not fully used in structure performance, and the value of reduction is within the requirements of gost 31938. the analysis of research allows estimating the service life of hcr in concrete for about 50 years, which corresponds to known data. the regularities of operation of hcr based on glass and carbon fibers in bent concrete beams are established. it is shown that the hcr can increase the rigidity of concrete beams by 15 % and crack resistance by 12 % in comparison with gfrp reinforcement, and the elastic modulus remains high (62 gpa) even after exposure to an alkaline environment. references [1] v. falikman, v. stepanova , a. bouchkin. russian experience in non-metallic composite reinforcement production and use. the fourth international fib congress 2014 (mumbai, india). improving performance of concrete structures. proceedings. university press. vol. ii, p. 784-786, 2014. [2] v. g. khozin. i. a. starovoitova, n. v. maisuradze, et al. nanomodification of polymer binders for structural composites. building materials 2:4-11, 2013. [3] a. d. pomogailo, a. s. rozenberg, i. e. upland. metals nanoparticles in polymers. khimiya, p. 672, 2000. [4] a. l. treniova. preparation of composite materials based on an epoxy oligomer and nanofillers, p. 12-39, 2008. [5] komar. composite reinforcement and masonry mesh manufactured. http://www.komarmatura.ru/. [6] y.-j. you, y.-h. park, h.-y. kim, et al. hybrid effect on tensile properties of frp rods with various material compositions. composite structures 80(1):117-22, 2007. https: //doi.org/10.1016/j.compstruct.2006.04.065. [7] russian patent no. 2287431. method of composite reinforcement manufacturing, 2006. [8] n. p. frolov. production technology of fiberglass reinforcement and some of its properties. concrete and reinforced concrete 9:5-8, 1965. [9] l. g. kuznetsova, y. s. cherkinsky, n. p. frolov, et al. improving of fiberglass reinforcement resistance. concrete and reinforced concrete, 3:30-31, 1973. [10] n. a. moshansky. on the fiberglass reinforcement resistance in concrete. concrete and reinforced concrete, 9:33-34, 1965. [11] n. k. rosenthal, g. v. chekhniy, a. r. bel’nik, et al. corrosion resistance of polymer composites in the alkaline environment of concrete. concrete and reinforced concrete, 3:20-23, 2002. [12] a. a. dalinkevich, k. z. gumargalieva, a. v. sukhanov, et al. aging kinetics of basalt and some glass fibers in an alkaline environment. plastics 12:23-26, 2012. 151 http://www.komarmatura.ru/ https://doi.org/10.1016/j.compstruct.2006.04.065 v. r. falikman, d. a. ilyin, v. f. stepanova acta polytechnica ctu proceedings [13] hughes brothers. glass fiber reinforced polymer rebar. hughes brothers, inc. 15, 1997. [14] w. ramm. report concerning the tests regard to the alkaline consistency of an anchoring of plastic reinforced with glass fiber concerning three-layered facade panels according to the deha-tm system, p. 32, 1992. [15] w. ramm. report concerning tests regarding the alkaline durability of an anchoring system out of reinforced glass fiber plastic for three-layered facade panels according to the deha-tm system, p. 43, 1993. 152 acta polytechnica ctu proceedings doi:10.14311/app.2019.25.0073 acta polytechnica ctu proceedings 25:73–78, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app mechanical and structural properties of collagen nanofribrous layers under simulated body conditions jitka říhováa, ∗, tomáš suchýb, lucie vištejnovác, lukáš hornýa, monika šupováb a czech technical university in prague, faculty of mechanical engineering, department of mechanics, biomechanics and mechatronics, technická 4, 166 07 prague 6, czech republic b czech academy of sciences, institute of rock structure and mechanics, v holešovičkách 94/41, 182 09 prague 8, czech republic c charles university, medical faculty in pilsen, biomedical center, alej svobody 1655/76, 323 00 pilsen, czech republic ∗ corresponding author: jitka.rihova@fs.cvut.cz abstract. the theme of this paper is the analysis of mechanical and structural properties of nanofibrous col under simulated body conditions and in the presence of osteoblasts and dermal fibroblasts. col were prepared by electrostatic spinning of 8 wt% collagen type i dispersion with 8 wt% (to col) of peg in phosphate buffer/ethanol solution (1/1 vol). the stability of col was enhanced by means of cross-linking with edc and nhs at a molar ratio of 4:1. col were exposed in culture medium for 21 days and human saos-2 human dermal fibroblasts and osteoblasts were cultured therein for 21 days as well. the cell culture on col was assessed by fluorescence microscopy and metabolic activity. then the metabolic activity of both cell types grown on col and ps were measured after 1, 7, 14 and 21 days using the alamar blue assay method. mechanical properties were determined using an tensile test. the influence of the cell activity on secondary structure of col was verified by ir spectroscopy. furthermore, the influence of cells on col was evaluated by sem. keywords: collagen, fibroblasts, osteoblasts, mechanical properties, structural properties, biodegradation. 1. introduction part of the tissue engineering is research and development of new materials which will be used in biomedicine. one of the directions is the development of biocompatible nanofibers which can be prepared, for example, from polymers. these nanostructures can provide temporary mechanical and structural support for cells and help restore damaged tissue. current research in the field are concentrated mostly on materials that are completely resorbable in the human body over time and their properties can simulate the function of the original tissue. understanding the factors that influence their degradation is crucial for the development of polymer degradable systems, especially in terms of degradation kinetics, changes in mechanical properties, and identification of degradation products [1, 2]. the term biodegradation has been known for degradation occurring in biological environment. in the context of biomedical applications, biodegradation can be defined as the gradual decay of a material driven by a specific biological activity [3, 4]. changes in the physico-chemical properties of materials that are exposed to body fluids are the result of chemical, physical, mechanical, and biological interactions between the material and the environment [5, 6]. 2. material and methods 2.1. nanolayer preparation collagen nanolayers (col) were prepared by electrostatic spinning (4spin, contipro, czech republic). the first step in the preparation was to mix the desired amount of collagen type i (calf skin, vup medical, czech republic) with polyethylene oxide (mw 900,000, sigma-aldrich, germany) and phosphate buffer (pbs, sigma-aldrich, germany). this mixture was incubated at 37 ◦c for 3 days. the hydrolysed collagen must be homogenized (10000 rpm, t25, ultra-turrax, ika, germany) to form a collagen dispersion, to which ethanol (penta, czech republic) is added. the electrospinning solution is prepared as 8 % w/w. collagen in a mixture of phosphate buffer and ethanol (1:1 by volume). the proportion of polyethylene oxide, as a auxiliary polymer which was completely washed after crosslinking of the nanofiller, was 8 % by weight (to collagen). after that col had to be cross-linked, which helps to increase its stability. the crosslinking was performed chemically with 95 % ethanol (penta, czech republic) with n-(3-dimethylaminopropyl)n-ethylcarbodiimide hydrochloride (edc, sigmaaldrich, germany) and n-hydroxysuccinimide (nhs, 73 http://dx.doi.org/10.14311/app.2019.25.0073 http://ojs.cvut.cz/ojs/index.php/app j. říhová, t. suchý, l. vištejnová et al. acta polytechnica ctu proceedings sigma-aldrich, germany). the edc / nhs molar ratio was 4:1. in this solution, the nanoparticles were soaked for 24 hours at 37 ◦c. finally, the col is rinsed with sodium phosphate (penta, czech republic) three times for 20 minutes. subsequently, the col were twice rinsed in distilled water, which is important for the complete washing of the polyethylene oxide. finally, washed col were frozen (−30 ◦c) and dried in a lyophilizer (virtris benchtop, usa) [7]. 2.2. exposure in culture medium and in the presence of cells for the exposure of col in simulated body conditions, dermal fibroblasts and osteoblasts (human line saos-2) cultivated on the surface of the layers were used. both cell types were first cultivated on 75 cm2 polystyrene culture bottles (techno plastic products, switzerland) in culture medium. the culture medium contained dulbecco’s modified eagle medium (dmem, thermo fisher scientific, usa), fetal bovine serum (fbs, thermo fisher scientific, usa), penicillin (100 u/ml), streptomycin (0.1 mg/ml), lglutamine, and 1 % non-essential amino acid (biosera, usa). cultivation took place at 37 ◦c and 5 % co2. on the day of the experiment, the dermal fibroblasts were sedimented with trypsin/edta and counted using a bürker cell. for fluorescence microscopy and for measuring metabolic activity, col were prepared in the form of a 0.3 cm circle, rinsed for 20 min in 1 % ethanol and 10 min in pbs. the samples were then placed in a 96 polystyrene panel (techno plastic products, switzerland) and immersed in 150 µm of culture medium until they were seeded with cells. for mechanical tests, 9 × 50 mm strips were rinsed for 20 minutes in 70 % ethanol and 10 minutes in pbs. 20 × 50 mm (spl life sciences, korea) culture chambers were each placed in two pieces and immersed in 4 ml of culture medium until they were fitted with cells. the dermal fibroblasts were diluted to a culture medium at a concentration of 30, 000 cells/ml and loaded onto collagen nanoparticles at a deposition density of 15, 000 cells/cm2 and a volume of 150 µl. the osteoblasts were diluted to a culture medium at 30, 000 cells/ml and loaded onto collagen materials at a loading density of 15, 000 cells/cm2 and a volume of 150 µl. both cell types were cultured for 21 days, when the measurements ranged on days 1, 7, 14 and 21. the culture medium was changed three times a week. 2.3. biological tests the cell culture on col were assessed by fluorescence microscopy and metabolic activity. first, the proper seeding by both cell types was evaluated. dermal fibroblasts and osteoblasts were dyed at 1, 7, 14 and 21 days after the start of the experiment. at each time, both cell types were cultured on col and on polystyrene (ps) stained with a 1 µg/ml solution of calcein am (thermo fisher scientific, usa) in culture medium for 30 min at 37 ◦c and 5 % co2. after the staining period, the samples were rinsed with pbs and transferred to the cover glass. clear fluorescence images of both types of cells cultured on col and ps were acquired using the olympus ix83 (olympus, switzerland) microscope at 488 nm excitation and 528 nm emission. images were processed using imagej (national institute of health, bethesda, usa). the metabolic activity of dermal fibroblasts and osteoblasts grown on col and ps were measured at 1, 7, 14 and 21 days after the start of the experiment using the alamar blue assay method. from each sample, culture medium was aspirated and 150 µl of new culture medium containing 10× diluted alamar blue (thermofisher scientific, usa) was added. samples were incubated for 2 hours at 37 ◦c and 5 % co2. subsequently, 100 µl of culture medium was pipetted into a 96 plate (nunc, denmark) and the fluorescence of the solution at 530 nm excitation and 590 nm emission was measured using a synergy ht (biotek, usa) spectrophotometer. from the measured values, the fluorescence of the culture medium itself was read and the values were expressed as the fluorescence ratio at day 7, 14 and 21 on fluorescence. values thus reflect the change in the metabolic activity of cultured cells on col and on ps over time. the experiment was performed in 6 replicates. 2.4. mechanical tests changes in mechanical properties were measured by a uniaxial tensile test in which the initial modulus of elasticity in the first and the last (fifth) load cycle and the ultimate tensile strength of the collagen layer were evaluated. the tensile test was carried out in the laboratory of biomechanics at the ctu in prague. measurements were carried out on a multi-axis test machine (zwick/roell, germany) with force sensors u9b by hbm with a range of ± 25 n and ± 250 n equipped with a non-contact optical extensometer. this extensometer, using the contrast marks on the samples, records the course of deformation during loading. the data was recorded at a frequency of 20 hz and stored on a computer. the position of the marks, the position of the clamping jaws and the force were measured. mechanical tests were performed for groups of 8 samples in medium, with osteoblasts and fibroblasts for each time interval. because of viscoelastic behavior of hydrated col, which also depends on the previous deformation, it is necessary to load and relieve the sample several times before performing the inter-strength test. estimation of the initial modulus of elasticity occurred both for the first load cycle, thanks to which we can determine the effect of dermal fibroblasts and osteoblasts on the surface layers of the material, as well as the last load cycle from which the effect of cells on the inner structure of the material can be estimated. from the measured data, graphs of stress σ strain � relationships were plotted. the values for these charts were calculated 74 vol. 25/2019 mechanical and structural properties of collagen nanofribrous from: σ = f s [mpa] , (1) � = ∆l l0 [−] , (2) where f is the force measured by the force sensors [n], s is the cross sectional area of the unloaded sample [mm2] (we consider the rectangular cross-section), ∆l is the elongation of the sample determined by the change of the contrast marks on the samples [mm] and l0 is the initial length of the sample [mm]. 2.5. structural changes structural changes were monitored by infrared spectroscopy and scanning electron microscopy. secondary structure of col was studied by infrared spectrometry (ftir), which allows study of its changes (denaturation) after various processes (isolation, crosslinking, sterilization). the aim of this analysis was to verify the possible effect of cell activity as well as exposure in the culture medium to the secondary collagen structure. the structure of col was studied using protégé 460 e.s.p. (thermo nicolet instruments, usa) by a method of attenuated total reflection atr (gladiatr, pike technologies, usa) with a diamond crystal. the resulting spectra were obtained from 128 scans with a resolution of 4 cm−1. belt areas (integral absorbance) were determined using omnic 7 software, which was also used for the deconvolution procedure of the amide bands i. the input data for band separation (number, position, half width, relative intensity and shape) were obtained using the fourier self-deconvolution procedure and further specified by the fitting procedure using the gaussian profile function. the individual col were measured at 10 different sites and the values of the waist areas after deconvolution were statistically evaluated. a sem analysis was used to basically assess the surface of collagen layers quanta 450 electron microscope (fei, usa). to assess the effect of cultivation on the surface of col the images (magnification 500× and 25, 000×) were taken at time intervals on day 1, day 7, day 14, and day 21. another reason of the sem analysis was to determine the fiber diameter before and after the crosslinking. to evaluate these data, the diameter of the fibers was evaluated manually by measuring their diameter (at magnification of 25, 000×). all samples were sputter-coated with a thin layer of gold (emissions k550x, quorum technologies, uk) prior to scanning. 3. results 3.1. biological test after the first day of culture, both dermal fibroblasts and osteoblasts were similar to the rounded shape of the cell on ps and on the col (see fig. 1). in the following days of cultivation (7th, 14th, 21st day) dermal fibroblasts on ps showed a typical elongated cell shape and covered most of the area planted. most cells were oriented in the same direction. in contrast, dermal fibroblasts on col did not show a one-way arrangement. the cells had a typical shape with elongated protrusions throughout the col space indicating their physiological comfort. on the 7th, 14th, and 21st days of cultivation, osteoblasts on ps showed a typical cuboid shape and covered most of the area planted. osteoblasts on col were more rounded, and the cells grew on multiple layers (see fig. 1). figure 1. dermal fibroblasts and osteoblasts cultivated on polystyrene and collagen were stained with calcein am solution, green color signifies metabolically active cells (scale 100 µm). compared to the first day, the metabolic activity of dermal fibroblasts on the col was 9 times higher (8.8±2.5) after 7 days of culture, and 5 times higher on ps (4.5 ± 0.6). this difference between col and ps was statistically significant (p < 0.05). after 14 and 21 days of culture, the metabolic activity of these cells on the col was 6 times higher (5.± 0.7 and 5.9 ± 1.4). on ps after 14 days 6 times and 21 days 5 times higher (5.6 ± 0.7 and 5.3 ± 0.4) versus the first day. these differences between col and ps were not statistically significant (p < 0.05) (see fig. 2). metabolic activity of osteoblasts was 10x higher (9.7 ± 2.9) on the col after 7 days of cultivation, and 11 × higher (11.0 ± 1.6) on the ps compared to the first day. this difference was not statistically significant (p < 0.05). after 14 days of cultivation, osteoblastic metabolic activity on col was 7 times higher (7.3 ± 0.5) and on ps nearly 13 times higher (12.7±1.8) versus day 1. after 21 days, their metabolic activity on col 10 × higher (10.2 ± 3.5) and on ps nearly 16 times higher (15.8 ± 0.9). differences between col and ps were statistically significant (p < 0.05) at both time points (see fig. 2). 75 j. říhová, t. suchý, l. vištejnová et al. acta polytechnica ctu proceedings figure 2. metabolic activity of dermal fibroblasts and osteoblasts grown on col and ps, * indicates statistically significant differences (p = 0.05, t-test) 3.2. mechanical properties from the measured data on samples exposed to medium without or with both cell types (n = 8) on the 1st, 7th, 14th and 21st days, stress-strain relationships were plotted in figs. 3 4. from the measured data, the initial modulus of elasticity was obtained as the slope of the tangent made to a stress-strain relationship on the initial linear part in the first and the last cycle. the ultimate tensile strength was determined as the maximum observed value of the nominal stress during the last loading cycle sample. the samples were labeled as follows: m for samples exposed only in medium, f for samples with dermal fibroblasts and s for samples with osteoblasts. the values of the initial elastic moduli in the first load cycle were not different for the samples exposed on the 1st and 21st day (see fig. 5). in both dermal fibroblasts and osteoblasts (see fig. 5), the values of the initial modulus of elasticity in the first loading cycle were decreased by 16 % and 36 %. a statistically significant difference in the last load cycle was observed only in osteoblast specimens (see fig. 5). for samples in the medium and fibroblast samples (see fig. 5), the difference in values was not statistically significant. for samples exposed in the medium, fibroblast samples and osteoblast samples, a statistically significant increase in the strength limit was recorded, by 12 % (see fig. 6). 3.3. structural properties 3.3.1. infrared spectroscopy collagenous materials were comparable from a spectroscopic point of view. the quartet of the ∼ 1205, 1240, 1280 (amide iii) and 1340 cm−1 quartets figure 3. stress-strain graphs for the first and last load cycles. for specimens exposed in 1. (m1) and 21 (m21), samples with fibroblasts (f1) and day 21 (f21) and samples with osteoblasts (s1) and day 21 (s21). figure 4. stress-strain graphs for the first and last load cycles. for specimens exposed in 1. (m1) and 21 (m21), samples with fibroblasts (f1) and day 21 (f21) and samples with osteoblasts (s1) and day 21 (s21). were a clear proof of the existence of a helical structure, in other words, the natural collagen structure was maintained at the appropriate level during the preparation of the layers by electrostatic spinning, to its fundamental violation. spectral decomposition of the amide band i made it possible to determine the ratio of the helical portion of collagen represented by the 1660 cm−1 band to the denatured structures based 76 vol. 25/2019 mechanical and structural properties of collagen nanofribrous figure 5. box plot comparing the initial tensile modulus in the first and the last cycle 1. day (m1) and 21. (m21) day of samples exposed in medium, 1. (f1) and 21. day (f21) of fibroblast, 1. (s1) and 21 (s21) of osteoblast samples, * indicates statistically significant differences (mann-whitney, 0.05) figure 6. box plot comparing the tensile strength limit of 1 (m1) and 21 (m21) of the day of samples exposed in medium, f1 and f21 of fibroblast samples, 1 (s1) and 21. day (s21) of osteoblast specimens, * indicates statistically significant differences (mannwhitney, 0.05) on the 1630 cm−1 bands and the 1615 cm−1 band. the 1660/(1630 + 1615) ratios were determined from measurements of 10 different sites of the collagen layers exposed to the culture medium (m1 and m21) and of the layers exposed to the cells. fibroblasts (f1 and f21) and osteoblasts (s1 and s21). the obtained data were statistically evaluated and compared. as can be seen, only the s1 and s21 pairs are statistically significant (see fig. 7). 3.3.2. electron microscopy the average fiber diameter of the sample before crosslinking was 261±49 nm. the average fiber diameter of the sample after crosslinking was 260 ± 35 nm. as can be seen from fig. 8 the fiber diameters before and after figure 7. box plot comparing ratio of helical and denatured areas of collagen (1660/(1630 + 1615)) 1. day (m1) and 21. (m21) day of samples exposed in medium, 1. (f1) and 21. day (f21) of fibroblast, 1. (s1) and 21 (s21) of osteoblast samples the cross-linking were not statistically significantly different. the mean fiber diameter of the sample did not changed during the exposure in the medium (202 ± 30 nm on the first day and 196±26 nm on the last day). the mean fiber diameter of the fibroblast samples after cross-linking the first day was 187 ± 33 nm, the last day 184 ± 32 nm. for osteoblast samples the mean fiber diameter of the after crosslinking was 198 ± 27 on the first day and 198 ± 25 on the last day. from fig. 9 it was obvious that the fiber diameter values were not statistically significant in the case of exposed samples. figure 8. a box plot comparing the fiber diameter before and after the crosslinking. fiber diameters are not statistically significant different. 4. conclusion during this work it was found that both fibroblasts and osteoblasts can influence the mechanical and structural properties of collagen layers by their activities. the activity of the cells has been demonstrated especially in the case of osteoblasts. in their presence there was a reduction in the modulus of tensile elasticity, which manifested itself in both the first and last load cycles. in the case of fibroblasts, this phenomenon was observed only in the first cycle, ie in a situation where the mechanical properties are more affected by the properties of the surface layers. in the case of an increase in the tensile strength after a three-week exposure, it was not shown that it was due to cell 77 j. říhová, t. suchý, l. vištejnová et al. acta polytechnica ctu proceedings figure 9. a box plot comparing the fiber diameter after first (m1) and 21 (m21) in medium, 1. (f1) and 21. (f21) days of fibroblast culture and 1. (s1) and 21. (f21) days of culture of fibroblasts and 1. (s1) and 21. (s21) days of cultivation with osteoblasts. fiber diameters are not statistically significant different. activity, since higher strengths were also achieved in samples exposed in the cell-free medium itself. higher activity of osteoblasts was further demonstrated by infrared spectroscopy, which showed an increase in the ratio of helical collagen to denatured structures. the presence of cells did not affect the diameter of the fibers of the collagen nanoparticle. fibroblasts and osteoblasts were succesfully plated and cultured on col. changes in the initial modulus of elasticity were observed most in the first load cycle, where both the fibroblast and the osteoblast samples showed statistically significant differences between the 1st and 21st day of the cultivation. for the last load cycle, statistically significant changes between 1st and 21st day of cultivation were observed only in samples with osteoblasts. on the other hand, ultimate tensile strenght increased in both cell culture media as well as in layers with cultured cells, thus the changes in ultimate tensile strength cannot be addressed only to cell activity. the infrared spectroscopy revealed the changes in helical structure of col only in osteoblast specimens. fiber diameters changes which were measured from images from electron microscopy were not observed. fibroblasts and osteoblasts can influence mechanical and structural properties of collagen nanoparticles. identification of cellular mechanisms responsible for the mechanical changes will be the aim of following research focused on the analysis of enzymatic activity and remodelation procedures of dermal fibroblasts and osteoblasts. acknowledgements the study was supported by project no. cz.02.1.01/0.0/0.0/16_019/0000787 "fighting infectious diseases", awarded by the ministry of education, youth and sports of the czech republic, financed from the european regional development fund. references [1] k. jahan, m. tabrizian. composite biopolymers for bone regeneration enhancement in bony defects. biomaterials science 4(1):25–39, 2016. doi:10.1039/c5bm00163c. [2] f. o’brien. biomaterials & scaffolds for tissue engineering. materials today 14(3):88–95, 2011. doi:10.1016/s1369-7021(11)70058-x. [3] h. azevedo, r. reis. understanding the enzymatic degradation of biodegradable polymers and strategies to control their degradation rate. crc press, boca raton, 2005. doi:0.1201/9780203491232.ch12. [4] l. ruiz-rubio, l. perez-alvarez, j. vilas-vilela. biodegradable shape-memory polymers. advanced structured materials 115:219–236, 2020. doi:10.1007/978-981-13-8574-2_10. [5] d. williams, s. zhong. biodeterioration/biodegradation of polymeric medical devices in situ. international biodeterioration and biodegradation 34(2):95–130, 1994. doi:10.1016/0964-8305(94)90002-7. [6] d. kytyr, p. zlamal, p. koudelka, et al. deformation analysis of gellan-gum based bone scaffold using on-the-fly tomography. materials and design 134:400–417, 2017. doi:10.1016/j.matdes.2017.08.036. [7] t. suchy, m. supova, e. klapkova, et al. the release kinetics, antimicrobial activity and cytocompatibility of differently prepared collagen/hydroxyapatite/vancomycin layers: microstructure vs. nanostructure. european journal of pharmaceutical sciences 100:219–229, 2017. doi:10.1016/j.ejps.2017.01.032. 78 https://doi.org/10.1039/c5bm00163c https://doi.org/10.1016/s1369-7021(11)70058-x https://doi.org/0.1201/9780203491232.ch12 https://doi.org/10.1007/978-981-13-8574-2_10 https://doi.org/10.1016/0964-8305(94)90002-7 https://doi.org/10.1016/j.matdes.2017.08.036 https://doi.org/10.1016/j.ejps.2017.01.032 acta polytechnica ctu proceedings 25:73–78, 2019 1 introduction 2 material and methods 2.1 nanolayer preparation 2.2 exposure in culture medium and in the presence of cells 2.3 biological tests 2.4 mechanical tests 2.5 structural changes 3 results 3.1 biological test 3.2 mechanical properties 3.3 structural properties 3.3.1 infrared spectroscopy 3.3.2 electron microscopy 4 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0123 acta polytechnica ctu proceedings 22:123–127, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app mechanical properties improvement of fiber reinforced concrete jan trejbala, ∗, václav nežerkaa, radim hlůžeka, zdeněk prošeka, b a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague, czech republic b czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: jan.trejbal@fsv.cvut.cz abstract. fiber reinforced concrete mechanical properties are limited due to low adhesion between polymer fibers and cement matrix. to ensure a strong interaction between the two materials, polypropylene fibers (d = 0.305 mm) were modified by an oxygen plasma treatment. the interface interaction was moreover activated using finely ground concrete recyclate, whose individual grains (1–64 µm) ensure an adhesion improvement in interfacial zones. the adhesion enhancement was verified by pull-out tests, when reference and modified fibers were pulled-out from cement matrix specimens. such obtained results were used as a crucial parameter to numerical simulations of bending tests of specimens (550 × 150 × 150 mm) with properties following fiber reinforced concrete. it was shown that samples reinforced with modified fibers and contained activating recyclate reached on higher residual bending strength then those with reference fibers. keywords: fiber reinforced concrete, macro fibers, polymer fibers, interfacial shear stress. 1. introduction fiber reinforced concrete (frc) has became popular at production of prefabricated concrete materials, shotcretes, and industrial high-loaded floors. such material is composed from polymeric macro-fibers (amount ca. up to 1 % vol. of whole mixture), cement, and aggregate [1, 2]. technical standards en 14845-1 and en 14845-2 describe frc as structural concrete reinforced with fibers having static effect and fulfilling requirements of en 14889-2. during three-point bending test of notched specimens 550 × 150 × 150 mm, such frc has to exhibit, besides, residual strength at least 1.5 mpa at crack mouth opening displacement (cmod) of 0.5 mm (corresponding deflection 0.47 mm). it is clear that such behavior differentiates the frcs from strain hardening or engineered composites, where strain-hardening response is required after reaching the elastic limit [1]. d. j. kim et al. explained that strength limit of a fibrous composite material (including frc) is a function of fibers volume, fibers length to diameter ratio, and the interfacial interaction between fiber surfaces and matrix. in the field of frcs, it means that increasing fibers amount weakens the cement matrix mechanical properties in the stage of elastics response during loading. it is therefore clear that fibers amount should be as small as possible. on the other side, once matrix limit of proportionality is reached and the matrix is damaged by the crack, fibers transfer the acting stress across the crack (crack bridging) and thus ensure macroscopic integrity of whole material. amount of stress transfered via fibers depends especially on their number and on adhesion between fiber surfaces and the cement matrix [3]. however, such adhesion is mostly too poor, especially between polymeric fibers and the cement matrix due to fibers smooth and chemically inert surfaces (related to cement matrix) [4, 5]. mechanical potential of fibers – tensile strength – is therefore unused. to avoid issues connected with the poor adhesion between the two materials, some researches have applied additional treatment of fibers in order to decrease their surface free energy and to increase their morphology, both ensuring improvement of bond to the matrix. for these purposes, several types of treatment may be employed, e.g. chemical (use of high alkali solutions) and physical (mechanical roughening) [6–8]. plasma modification has shown to be a promising technology, combining both the chemical (etching) and physical (roughening caused by a ion bombardment) treatment, as proven by li et al. from the early 1990s [9, 10] and many other researches later, e.g. [11– 14]. it is also worth noting that such treatment has been extended through many industrial fields over the past few years, especially for the surface treatment (roughening, activating, cleaning) of polymeric materials [15]. therefore, there is no obstacle to apply such technology during surface treatment of the fibers. although a benefit of fiber surface treatment was proven from the perspective of “surface science” many times, this was not achieved from the practical point of view, including the field of frcs. to connect theoretical findings with praxis of civil engineering, we studied an influence of plasma modified fibers on 123 http://dx.doi.org/10.14311/app.2019.22.0123 http://ojs.cvut.cz/ojs/index.php/app j. trejbal, v. nežerka, r. hlůžek, z. prošek acta polytechnica ctu proceedings mechanical properties of frc samples using numerical simulations, following en 14845-1. 2. interaction between fiber and matrix post-cracking response of frc is influenced by behavior of one fiber that is pulled out from the matrix. this phenomenon was described by ch. li et al. and c. redon et al. [16, 17]. the behavior is divided into two stages, the first describes chemical interaction between the two materials (pdeb), while the second mechanical interaction activated by fiber movement out of the matrix (ppull), concretely: pdeb = √ π2efd 3 f 2 (τ0u + gd) (1) and ppull = πdfτ0  1 + β ( u − ( 2τ0l2e ef df + √ 8gdl2e ef df )) df   × [ le − u + ( 2τ0l2e efdf + √ 8gdl2e efdf )] , (2) and τ0, as a frictional stress after sudden drop following the peak pull-out load, by: τ0 = ppull πdfle (3) where ef is young’s modulus of elasticity of fibers; df, fibers diameter; τ0, interfacial shears tress between fiber surface and matrix; u, fiber free-end displacement; gd, interfacial bond strength; β, a shear retention factor, parameter describing slip softening / hardening behavior; le, fiber embedded length. 3. fibers and their treatment 3.1. polymeric fibers polymeric macro-fibers were used for all experiments described below. their geometrical and mechanical properties were as follows: material, polypropylene (pp); morphology, smooth; length, 60 mm; diameter, 305 µm; density, 900 kg/m3; young’s modulus of elasticity, 6.1 gpa; tensile strength, 440 mpa; elongation, 8 %. mechanical properties have been determined experimentally, as reported in [18]. 3.2. fiber treatment two fiber types were used: reference (further marked as r) and plasma treated (p30 and p120 according to treatment duration). plasma treatment was executed using tesla vt214 device. treatment parameters were: plasma, cold; gas, oxygen; power of rf source, 100 w; gas pressure, 20 pa; treatment duration, 30 and 120 seconds. name cement recyclate w/c [wt. %] [wt. %] [-] ref 100 0 0.40 rec 70 30 0.41 table 1. weight proportions of cement matrices. 4. pull-out tests two matrices were used to carry out pull-out tests. the reference matrix (ref) was made from portland cement cem i 42.5r and the modified matrix (rec) contained 30 wt. % concrete recyclate as a substitution for cement at the form of finely ground powder. the recyclate was used to fill interfacial zones between fiber surfaces and surrounding matrix and thus to ensure the adhesion between the two materials with individual grains. grain size differed between 1–64 µm, as measured by blain’s method. matrices compositions are summarized in table 1. specimens made from such matrices had dimensions equal to 25 × 20 × 30 mm, contained a single fiber in their centerline (fiber embedded length was equal to samples height – 30 mm). results of this experiment were used as basic input values for numerical modeling of frc bending tests. the whole experiment was carried out using loading frame veb tiw rauenstein fp100. the specimen was anchored by its matrix body to a static part of the frame, while the single fiber, protruded from the specimen body, was caught by a moving frame part. the experiment was displacement controlled at the constant rate of 3 mm/min, finished after reaching to 4.5 mm of fiber-free end displacement (only r and p30 were further used). results from pull-out tests – dependence between fiber free-end displacement and force resisting to that – are summarized in figure 1. it is clear from these results that the maximal force recorded during pullout reference fiber (r) from the reference matrix (ref) slightly overcame 10 n, while in the case of 30 seconds plasma modified fibers (p30) and matrix containing concrete recyclate (rec), the force reached on more than 14 n. despite of the assumptions, fibers exposed to plasma for 120 seconds (p120) exhibited adhesion to the matrix worse than these modified for 30 seconds. this could be caused by their diameter reduction due to too long treatment. therefore, such fibers were not used for numerical simulations described below. according to equation (4), shear stresses were calculated from thus obtained results. it was calculated that τ0,r = 3.24 · 105 pa andτ0,p30 = 4.56 · 105 pa in the case of reference and 30 seconds modified fibers, respectively. 5. numerical simulations numerical simulations followed procedure of threepoint bending test described in technical standard 124 vol. 22/2019 mechanical properties improvement of fiber reinforced concrete fo rc e [n ] 0 2 4 6 8 10 12 14 16 0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 fiber free-end displacement [mm] r p30 p120 figure 1. pull-out behavior of reference and modified fibers. figure 2. geometry of frc specimen (top) and mesh of computational model (bottom). en 14845-2. as already mentioned in introduction, notched frc specimen 550 × 150 × 150 mm, containing 0.5 vol.% of fibers, has to exhibit at least 1.5 mpa of bending strength in post-cracking phase during midspan deflection of 0.47 mm. to avoid lengthy experimental testing, numerical simulation was employed. shcc material model [19, 20], suitable also for frcs, was applied. mash of the 2d model, counting 3080 of predominantly triangle linear elements, was created in salome software [21]. non-linear numerical analysis was conducted using oofem software [22], proceeded in 800 steps. numerical experiment was controlled by displacement. solution was searched by newton-raphson’s method. the stiffness matrix was compiled in 2d plain-stress preposition. geometry of the specimen and mesh of finite elements are shown in figure 2. two calculations were done; the first contained reference, while the second one 30 seconds plasma modified fibers. mechanical properties of matrix was set to correspond common concrete. all parameters set to numerical model were as follows (reference / modified fibers): e young’s modulus of elasticity of matrix, 20 / 20 gpa; ν poisson’s ratio of matrix, 0.2 / 0.2; gf fracture energy of matrix, 5.0 / 5.0 n/m; ft tensile strength of matrix, 2.5 / 2.5 mpa; softtype a parameter describing post-peak behavior, 3 / 3: hordijk’s softening; sheartype a parameter describing shear stiffness of cracked material, 1 / 1: constant shear retention; shearstrengthtype a parameter limiting the magnitude of resulting shear stress acting on crack plane; 1 / 1: the threshold is set to the value of the tensile strength; vf fiber volume ratio, 0.005 / 0.005; df fiber diameter, 0.305 / 0.305 mm; df fiber length, 60.0 / 60.0 mm; ef young’s modulus of elasticity of fibers, 6.1 / 6.1 gpa; gf shear modulus of fibers, 1.0 / 1.0 gpa; τ0 frictional shear stress between the fiber and the matrix during debonding, 0.324 / 0.456 mpa; f snubbing coefficient; 0.5 / 0.5; kf fiber cross-section shape correction factor, 0.9 / 0.9; fsstype a class describing type of fiber bond shear strength, 0 / 0: constant shear strength; fibertype class of reinforcing fibers, 2 / 2: short randomly oriented fibers; ncracks maximal number of cracks, 2 / 2; m exponent related to fiber unloading, 1 / 1; fibreactivationopening, 10−6 / 10−6; dw0 lower bond allowing to smoothen the traction-separation law for fibers, 10−7 / 10−7; dw1 upper bond allowing to smoothen the traction-separation law for fibers, 10−7 / 10−7. 6. results it was found from numerical simulations that residual strength of frc specimen reinforced with plasma modified fibers (p30) at amount of 0.5 vol.% tightly exceeded 1.5 mpa at cmod of 0.47 mm. in the same stage, the samples containing reference fibers exhibited only ca. 1.3 mpa. results from both simulations are imagined in figure 3, where the green line highlights the minimal bending strength required by en 148452. it is obvious from these results that only frc containing modified fibers fulfilled these requirements, so this can be considered as structural. these simulations also revealed that the behavior of both specimens was practically identical in phases 125 j. trejbal, v. nežerka, r. hlůžek, z. prošek acta polytechnica ctu proceedings fiber free-end displacement [mm] 0 0 0.1 0.2 0.3 0.4 0.5 0.6 1.0 1.5 2.0 2.5 3.0 0.5 st re ss [m pa ] midspan deflection [mm] r p30 figure 3. normal tension as a function of midspan displacement. until the fibers have not been activated yet. after, fibers bridged the opening crack, transfered acting stress and thus ensured macroscopic integrity of specimens. based on the specimens post-cracking behavior, it is clear that the adhesion between modified fibers and the matrix was increased than in case of reference fibers. this finding proves that interfacial shear stress between the two materials plays an important role from the mechanical response point of view. 7. conclusions this work deals with searching of mechanical behavior of fiber reinforced concrete using numerical simulations of three-point bending tests. reference and plasma modified polypropylene fibers (d=0.305) were used as reinforcement. adhesion between these fibers and two types of the cement matrix was examined employing pull-out test. the purpose of the simulation was to find residual bending strength of specimens, following relevant technical standards. finding were as follows: • adhesion between 30 seconds oxygen plasma treated fibers and the cement matrix containing concrete recyclate was higher by approx. 15 % then in the case of reference fibers and reference cement matrix. • 120 seconds lasting plasma treatment did not bring any benefits in terms of adhesion improvement. conversely, such treated fibers showed worse adhesion to the matrix than those 30 seconds exposed to plasma. this was probably caused by their diameter reduction as a consequence of too intensive ion bombardment. • numerical simulations revealed that residual strength of reference frc at the midspsan deflection of 0.47 mm was less than 1.5 mpa, so this material did not meet requirements of technical standard en 14845-2. on the other side, if plasma treated fibers and the cement matrix containing 30 wt.% of concrete recyclate were used, residual strength overcame minimal 1.5 mpa. acknowledgements this work was financially supported by czech technical university in prague – sgs project sgs16/201/ohk1/3t/11 and sgs18/037/ohk1/1t/11, and by the czech science foundation research project 17-06771s. special thanks to tereza horová for research assistance. references [1] m. prisco, g. plizzari, l. vandewalle. fibre reinforced concrete: new design perspectives. materials and structures 42(9):1261–1281, 2009. 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[22] b. patzák. oofem – an object-oriented simulation tool for advanced modeling of materials and structures. acta polytechnica 6(6):59–66, 2012. 127 http://dx.doi.org/10.1016/s1065-7355(97)90001-8 http://dx.doi.org/10.1016/j.jmatprotec.2009.02.015 http://dx.doi.org/10.1016/j.cemconcomp.2018.03.010 http://dx.doi.org/10.1016/j.surfcoat.2013.10.008 http://dx.doi.org/10.1016/j.polymdegradstab.2012.08.001 http://dx.doi.org/10.1061/(asce)0733-9399(1992)118:11(2246) http://dx.doi.org/10.1061/(asce)0899-1561(2001)13:6(399) acta polytechnica ctu proceedings 22:123–127, 2019 1 introduction 2 interaction between fiber and matrix 3 fibers and their treatment 3.1 polymeric fibers 3.2 fiber treatment 4 pull-out tests 5 numerical simulations 6 results 7 conclusions acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0057 acta polytechnica 27(0):57–61, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app laser-induced surface acoustic waves for thin film characterization radim kudělkaa, lukáš václaveka, jan tomáštíkb, sabina malecováa, radim čtvrtlíkc,∗ a joint laboratory of optics, palacky university olomouc, 17. listopadu 50a, 772 07, olomouc, czech republic b institute of physics of the czech academy of sciences, joint laboratory of optics of palacky university and institute of physics as cr, 17. listopadu 12, cz-771 46, olomouc, czech republic c regional centre of advanced technologies and materials, joint laboratory of optics of palacky university and institute of physics of the czech academy of sciences, faculty of science, palacky university, 17. listopadu 12, 77146 olomouc, czech republic ∗ corresponding author: radim.ctvrtlik@upol.cz abstract. knowledge of mechanical properties of thin films is essential for most of their applications. however, their determination can be problematic for very thin films. law (laser-induced acoustic waves) is a combined acousto-optic method capable of measuring films with thickness from few nanometers. it utilizes ultrasound surface waves which are excited via short laser pulses and detected by a pvdf foil. properties such as young’s modulus, poisson’s ratio and density of both the film and the substrate as well as film thickness can be explored.results from the law method are successfully compared with nanoindentation for young’s modulus evaluation and with optical method for film thickness evaluation and also with literature data. application of law for anisotropy mapping of materials with cubic crystallographic lattice is also demonstrated. keywords: anisotropy, laser-induced surface acoustic waves, nanoindentation, thin films, young’s modulus. 1. introduction thin films and coatings have become an essential part of our everyday lives and knowledge of their mechanical properties and durability directly affects their lifetime regardless of their primary role. thin films are often exposed to external stresses during their fabrication or service life. therefore, their reliable testing is essential for both their research and development as well as industrial production. nanoindentation has become the standard method for determination of local mechanical properties, but it has its inherent limitations [1]. mostly the thickness of the film is the main limiting factor, because if the indenter penetrates to a greater depth than 10 % of the film (commonly used rule of thumb), then substrate may affect the hardness values considerably [2–4]. determining the elastic modulus of very thin films is even more problematic because the substrate and the film behave like two springs in series hence the elastic modulus of the film is always affected by the substrate [5]. this means that the indentation of films with thickness under 50 nm is very challenging [6]. however, mechanical properties of very thin films can still be explored using the methods based on analysis of the dispersion curves of rayleigh surface acoustic waves. one of such methods is the law (laser-induced acoustic wave) technique [7, 8]. it is a combined acousto-optic method developed for the assessment of material characteristics of surfaces (young’s modulus, poisson’s ratio, density, film thickness) based on the generation of surface acoustic waves and their detection. the propagation of these waves in the material depends on the elastic properties of both the thin film and the substrate. in order to demonstrate the possibilities and strengths of the law, the representative results from thin films and single crystals are presented. in case of thin films the results are compared with the results from other methods such as nanoindentation for young’s modulus evaluation or optical methods for film thickness characterization. besides, the limit of the film thickness is explored, as the ability to characterize very thin films is one of the main advantages of law over other methods. 1.1. principle of law short laser pulses from nitrogen laser are absorbed by the material and cause generation of a wide frequency spectrum (typically in the range from units to hundreds of mhz) of thermally activated surface acoustic waves (saw). those waves are propagating from the point of incidence and mostly in the surface layers of the sample as their amplitude decays exponentially with the increase of the distance from the surface [7] as shown in the figure 1. the penetration depth hpen of the saw can be estimated by their wavelength (hpen ≈ λ). this 57 http://dx.doi.org/10.14311/app.2020.27.0057 http://ojs.cvut.cz/ojs/index.php/app r. kudělka, l. václavek, j. tomáštík et al. acta polytechnica amplitude de pt h film substrate low frequency, high wavelength (a) low frequency, high wavelength figure 1. propagation and amplitude-decay of saw with a) low frequency and b) high frequency means that the penetration depth depends on the frequency f(λ = v/f ). the greater the frequency the lower the penetration depth. because most of the wave energy is concentrated on the surface, saw are very sensitive to very thin films with thicknesses from few nanometers (even when they are much thinner than the penetration depth of saw) [8]. the phase velocity v of saw in a non-coated homogeneous and isotropic material is independent on frequency and can be described as: ν = 0.87 + 1.12ν 1 + ν √ e 2ρ(1 + ν) , (1) where ν is poisson’s ratio, ρ is material density and e is young’s modulus. this means that phase velocity is directly linked to the elastic modulus, poisson’s ratio and density of the material. however, situation is different in coated materials, where the penetration depth depends on frequency so the phase velocity of the saw must as well. this phenomenon is called the dispersion. in case of a substrate with low phase velocity coated with a thin film with a high phase velocity, then the phase velocity for the film/substrate system increases with frequency. for explanation let’s assume infinite substrate and saw propagating across the surface. theoretically for the wave with frequency approaching zero, the penetration depth would be approaching infinity. that in turn means that the contribution from the thin film would be negligible. on the contrary, if the saw would have infinitively high frequency, then the wave would be restricted only to the very top of the sample, in other words the wave would propagate only in the film according to the eq. (1). the precise relation between the frequency f and phase velocity v of the saw on coated materials depends on elastic properties and densities of both the substrate and the thin film as well as on the thickness of the film. this means that if the dispersion curve and some of the variables are known, the rest of them can be calculated. in order to get the dispersion curve the saw must be detected in various known distances from the point of incidence using a pvdf foil which transforms the mechanical vibrations into electric signal, see principal scheme of law method (instrument) in figure 2. the dispersion curve is then calculated using the following equation: ν(f) = 2πf (x2 − x1) φ2(f) − φ1(f) , (2) where x1 and x2 are two different distances between the sensor and the point of incidence, φ1(f) and φ2(f) are the phase values of the surface wave frequency f. the whole experimental set-up can be seen in figure 2. substrate laser oscilloscope saw v pvdf foil substrate figure 2. schematic representation of the law set-up. 1.2. experimental details the experiments were performed on a laser acoustic test system lawave (fraunhofer institute, dresden, germany). acoustic waves with frequency ranging from ∼ 1 mhz up to ∼ 250 mhz were generated using a 337 nm n2 laser with a 0.5 ns pulse duration and 58 vol. 27/2020 laser-induced surface acoustic waves for thin film characterization 35 93 142 163 34 90 136 156 0 50 100 150 200 1 min 60 min 120 min 180 min f il m t h ic k n e ss [ n m ] law optical figure 3. film thickness of samples from dry thermal growth measured by law compared to results from optical method. 400 µj pulse energy. the signal from the pvdf foil was averaged from 64 measurements for each of the 5-7 measuring points along the 15 mm and 20 mm measuring path for glass and silicon substrates respectively. all measurements were done at a room temperature. 2. results and discussion 2.1. thin film characterization – film thickness evaluation silicon dioxide thin films on si(100) were used for the film thickness characterization. it is well known that high quality sio2 films are prepared by dry thermal growth which is considerably slower than for example wet thermal growth but yields high quality films – the oxide film is more uniform and its density is higher. this allows for more precise and reliable comparison of law and other methods as these sio2 films represent an ideal model system. the si substrates were oxidized at 1000 ◦c in quartz tube for 1, 60, 120 and 180 minutes, the heating rate was 20 ◦c/min. the resulting film thickness can be seen in figure 3. the optical reference measurement of the film thickness was performed on spectral reflectometer calibrated on polished silicon wafer with known profile of spectral reflectance. results from both methods are similar with the results from law being slightly higher in all cases. the highest percentage difference (4.6%) is for the thinnest sample. the standard deviation for law is very small. thus, it may be concluded that law gives reasonable results for film thickness if young’s modulus and density are uniform and known. the native sio2 oxide layer (grown at room temperature) was measured as well in order to find out how accurate can law be for very thin films. its thickness, when kept in dry pure oxygen, is around ∼1 nm [9], but thickness around 1.5-2.5 nm can be expected when exposed to air depending on the temperature and humidity with 2.1 nm given in [9]. since law is very sensitive to the top surface conditions as well as presence of (sub)surface flaws [10] and any structural imperfections, different si(100) wafers of the same type and doping 65 67 65 66 101 85 67 68 0 20 40 60 80 100 120 34 nm 90 nm 136 nm 156 nm y o u n g ’s m o d u lu s [g p a ] law nanoindentation figure 4. young’s modulus of sio2 thin films with different thickness measured by law and nanoindentation. were explored (university wafer, usa). this was confirmed by a strong variance in measured thickness of native sio2 film between different wafers. the native sio2 thickness of (3.2 ± 0.5) nm was obtained after the exclusion of the most extreme values that might be among others related to the sub-surface damage. these values are slightly higher in comparison with the literature data. this overestimation might be caused by the silicon suboxide interface layer formed between the si and sio2, with thickness estimated to be ∼0.35 nm [11]. the interface behaves like a second thin film and might increase the resulting film thickness since in our fit only single film system with characteristics of sio2 is assumed. this means that even few nm thick films may be explored via law method provided the dispersion caused by the film can be easily seen and substrate of superb quality, with no surface and sub-surface damage, is used and effects of interface layers are included in the calculation. 2.2. thin film characterization – young’s modulus evaluation in order to explore the inherent limits of depth sensing indentation and law the sio2 films with different thicknesses were measured by both methods. the high quality sio2 films with thickness ranging from 34 to 156 nm grown on si(100) mentioned in previous chapter were used. nanoindentation experiments were performed using the fully calibrated nanotest instrument equipped with a brand new berkovich indenter. at least 15 measurements for each film were analyzed by standard oliver-pharr method. the indentation modulus e values were calculated using sio2 poisson’s ratio of 0.17 and considering the correction for diamond elastic deformation (poisson’s ratio of 0.07 and young’s modulus of 1141 gpa). it should be noted that indentation modulus can be considered as young’s modulus as long as pile-up and sink-in effects are absent [12]. comparison of the results can be seen in figure 4. figure 4 shows that the values obtained from law and nanoindentation are similar for the 156 nm and 136 nm thick films because nanoindentation was not strongly affected by the substrate. however, for the 90 nm film the indenter penetrated up to 25 % of 59 r. kudělka, l. václavek, j. tomáštík et al. acta polytechnica 120 80 40 160 120 200 80 40 0 200 160 y o u n g ’s m o d u lu s [ g p a ] (a) p o is s o n ’s r a ti o 0.00 0.30 0.25 0.20 0.15 0.10 0.05 0.10 0.15 0.20 0.25 0.30 0.05 (b) figure 5. si(100) anisotropy (a) young’s modulus (b) poisson’s ratio. the film and the results are strongly affected by the substrate which has higher young’s modulus than the film. the 34 nm thick film shows even stronger substrate effect as 59 % of the film was penetrated and the resulting value of young’s modulus is somewhere in the middle between the value of the film and the value of the substrate. this clearly shows the unambiguous advantage of the law as the results obtained from law are constant despite the changing film thickness as opposed to nanoindentation. this is crucial for the characterization of very thin films. law can be affected by the film thickness as well, but only in the case of much thinner films as shown in the previous chapter reporting on characterization of native sio2 layer of ∼2 nm thickness. comparison in figure 4 shows that the elastic modulus values obtained by nanoindentation are almost identical to those measured by law as long as the film thickness is high enough. in this case for the films thinner than 100 nm the unavoidable substrate effect leads to overestimation of the elastic modulus. for such thin films law can be viable alternative. in order to present the robustness and universal use of law other films were also explored. young’s modulus was evaluated for aluminum film on glass, and titanium and chromium on silicon. the thickness of all three films was around 100 nm. values obtained by law are compared with the values given by other authors [13–17] and in case of al and ti also with the results from nanoindentation. table 1 shows that, the values of young’s modulus obtained by law correspond to those obtained from nanoindentation and are within the range reported in literature. 2.3. silicon anisotropy mapping single-crystal silicon is a well-known material with interesting properties and wide range of applications [18] that exhibits a strong anisotropy. therefore, it is an ideal object for testing the law ability to explore material anisotropy. material young’s modulus (gpa)law reference indentation al 63.5 55-81 69.6 ti 116.1 102-123 111.5 cr 204.4 185-215 – table 1. comparison of young’s modulus values of al, ti, cr thin film measured by law, nanoindentation and literature data in general, elastic behavior of cubic crystals may be fully described by three elastic stiffness constants c11, c12 and c44 (for silicon at room temperature c11 = 165.6–165.7, c12 = 63.9, c44 = 79.5–79.6 gpa [18, 19]). young’s modulus and poisson’s ratio can be calculated in any direction from those constants. for example, si(100) has young’s modulus ranging from 130.2 to 168.9 gpa [20], the minimal value is for the [100] direction and can be calculated as: e[100] = c11 − 2 c12 c11 + c12 c12. (3) in this work, the elastic stiffness constants were measured in various directions on several si (100) wafers. the resulting average values were c11 = 165.4, c12 = 63.5 and c44 = 79.6 gpa that is in excellent agreement with literature values listed above. using these constant it is possible to plot a polar dependence of young’s modulus and poisson’s ratio. figure 5 clearly shows the strong anisotropy of silicon elastic constants. using law, anisotropy of any cubic crystals can be explored and elastic constants can be calculated for any other plane orientations. 3. conclusions law is a combined acousto-optic method developed for assessment of material characteristics of surfaces and can be used to obtain values of young’s modulus, poisson’s ratio and density for both the film as well as 60 vol. 27/2020 laser-induced surface acoustic waves for thin film characterization the substrate. it can also be used to obtain the values of elastic stiffness constants for anisotropic materials with cubic symmetry. films’ thicknesses measured via law are very close to those from optical thickness measurements if input parameters are precisely known. young’s modulus measurements by law and by nanoindentation yielded same results for films with thickness over 100 nm, however as the indentation depth / film thickness ratio increased the data from nanoindentation started to be affected by substrate as the predicted by the commonly used 10 % rule of thumb (indenter should not penetrate deeper than to the 10 % of the film). acknowledgements the authors gratefully acknowledge the support from the project iga_prf_2019_008 of palacky university and the project th03020245 of the technology agency of the czech republic and the operational program research, development and education, project no. cz.02.1.01/0.0/0.0/17_049/0008422 of the ministry of education, youth and sports of the czech republic. references [1] w. oliver, g. pharr. measurement of hardness and elastic modulus by instrumented indentation: advances in understanding and refinements to methodology. journal of materials research 19(1):3–20, 2004. doi:10.1557/jmr.2004.19.1.3. 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[20] m. a. hopcroft, w. d. nix, t. w. kenny. what is the young’s modulus of silicon? journal of microelectromechanical systems 19(2):229–238, 2010. doi:10.1109/jmems.2009.2039697. 61 https://doi.org/10.1557/jmr.2004.19.1.3 https://doi.org/10.1016/s0257-8972(97)00522-7 https://doi.org/10.1016/0020-7683(88)90091-1 https://doi.org/10.1557/jmr.2009.0102 https://doi.org/10.1007/978-1-4757-5943-3 https://doi.org/10.1557/jmr.1999.0299 https://doi.org/10.1016/s0257-8972(01)01664-4 https://doi.org/10.1016/s0257-8972(96)03147-7 https://doi.org/10.1007/s12633-015-9366-8 https://doi.org/10.1016/j.mee.2006.11.001 https://doi.org/10.1116/1.2218865 https://doi.org/10.3403/bseniso14577 https://doi.org/10.3390/coatings7090143 https://doi.org/10.1016/0040-6090(95)06990-9 https://doi.org/https://doi.org/10.1016/j.tsf.2005.02.024 https://doi.org/10.1016/j.surfcoat.2012.02.039 https://doi.org/10.1016/s0040-6090(96)09029-3 https://doi.org/10.1107/s1600577514004962 https://doi.org/10.1007/978-3-642-59497-7_157 https://doi.org/10.1109/jmems.2009.2039697 acta polytechnica 27(0):57–61, 2020 1 introduction 1.1 principle of law 1.2 experimental details 2 results and discussion 2.1 thin film characterization – film thickness evaluation 2.2 thin film characterization – young's modulus evaluation 2.3 silicon anisotropy mapping 3 conclusions acknowledgements references doi:10.14311/app.2021.30.0069 acta polytechnica ctu proceedings 30:69–75, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app comparing the hoek-brown and mohr-coulomb failure criteria in fem analysis tereza poklopová∗, veronika pavelcová, michal šejnoha czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: tereza.poklopova@fsv.cvut.cz abstract. this paper revisits the issue of a potential substitutions of the hoek-brown failure model by the standard mohr-coulomb model in the stability analysis of rock masses. the derivation of equivalent shear strength parameters of the mohr-coulomb proposed by hoek et al. [1] is addressed with emphases on the suitable range of stresses for which the equivalence of the two failure criteria applies. to that end, a simple numerical analysis of the oedometric test is carried out. it is seen that a correct choice of the upper limit of the minimum compressive principal stress is crucial for the mohr-coulomb model to provide predictions comparable to the hoek-brown model. this issue is addressed next in the light of the solution of slope stability problem. all the presented results were derived with the help of the geo5 fem finite element software [2]. keywords: equivalent shear strength, finite element method, hoek-brown criterion, nonlinear yield function, rock mechanics. 1. introduction application of the mohr-coulomb (mc) model is a common approach when addressing the stability of soil slopes. although for a rock mass the hoek-brown (hb) failure criterion [1, 3–5] proved more appropriate, the use of the mohr-coulomb model in the stability assessment of rock slopes is still quite popular in practice. this can be attributed to the fact that the factor of safety can easily be defined in terms of the basic shear strength parameters such as the cohesion c and the angle of internal friction ϕ. this is supported by the possibility of estimating these parameters through simple empirical formulas on the basis of hb model parameters [1]. this, however, requires the knowledge of the range of stresses for which the linear mc model provides results comparable to those predicted by the nonlinear hb model. so it is perhaps more straightforward to proceed in the footsteps of benz et al. [6] who introduced a simple modification to the original hb model, which brings the two models to the same footing and allows for a standard strength reduction approach to estimate the factor of safety in the light of the hoek-brown model. both issues are examined in this paper. 2. hoek-brown model the hoek-brown model was first introduced by hoek [3] in application to the analysis of an underground excavation of intact rock masses. later in [4] this criterion was extended to weak rock masses by adjusting the material constants according to a geological quality. the geological strength index (gsi) is used as a qualitative classification number. by implementing the influence of prior excavation process, the current form of the hb model covers the whole range of rock masses. the generalized hoek-brown failure criterion is defined in terms of principal stresses as σ3 = σ1 − σci ! mb −σ1 σci + s "a , (1) where σ1 and σ3 are the maximum and minimum effective principal stresses, respectively1. the material constants mb, s and a follow from empirical equations as mb = mi exp gsi − 100 28 − 14d , (2) s = exp gsi − 100 9 − 3d , (3) a = 1 2 + 1 6 ! exp −gsi 15 − exp −20 3 " , (4) where the disturbance factor d account for prior excavations and ranges from 0 (undisturbed rock) to 1 (disturbed by production blasting). the uniaxial compressive strength σci and the model parameter mi are considered for an intact rock. the rock mass stiffness is assumed constant and independent of the current stresses. following [1], the modulus of elasticity is estimated by em(gpa) = ! 1 − d 2 " # σci 100 10((gsi−10)/40), (5) em(gpa) = ! 1 − d 2 " 10((gsi−10)/40). (6) 1unlike typical geotechnical notation, the tensile stresses are assumed positive. 69 http://dx.doi.org/10.14311/app.2021.30.0069 http://ojs.cvut.cz/ojs/index.php/app t. poklopová, v. pavelcová, m. šejnoha acta polytechnica ctu proceedings equation (5) is called when σci ≤ 100 mpa and suggests a rapid decrease of the modulus with a decreasing uniaxial compressive strength. equation (6) is then used when σci > 100 mpa. the above equations were exploited in this paper regardless of their updated version presented in [5]. 3. analogy with the mohr-coulomb criterion in the finite element analysis the failure criterion has a meaning of a yield function bounding the elastic response. the geo5 fem program assumes the stressstrain relation to be elastic perfectly plastic. the hoek-brown yield function then follows from eq. (1) in the form f (σ1, σ3) = σ1 − σ3 − σci ! mb −σ1 σci + s "a . (7) when presented in the principal stress space, the hoek-brown yield function is similar to the mc model independent of the intermediate principal stress σ2. in case of plastic yielding, the stress return mapping is driven by a non-associated flow rule with the plastic potential function given by g(σ1, σ3) = 1 + sin ψ 1 − sin ψ σ1 − σ3, (8) where ψ is the dilation angle and for the assumed perfect plasticity it remains constant. similar to mc model, the hb model plots as irregular hexagon, but unlike the mc model it experiences a nonlinear variation in the deviatoric plane, see also the projection into σ1-σ3 plane displayed in fig. 1. figure 1. hoek-brown yield function plotted in plane of maximum and minimum principal stresses. 3.1. equivalent shear strength parameters the simplicity of the mc model and its prevailing popularity among design engineers promote its use in applications concerning the stability issues of a rock mass. the similarity of the mc and hb models was examined by hoek et al. in [1, 4] in order to derive the shear strength parameters ϕ, c controlling the mc failure criterion as, compare with eq. (1), σ3 = −σcm + kσ1, (9) where σcm is the uniaxial compressive strength of the rock mass and k is the slope of the mc plot in the σ3 and σ1 principle stress space, see fig. 2. effective values of the shear strength parameters c, ϕ are then given by c = σcm 2 √ k , sin ϕ = k − 1 k + 1 . (10) comparing eqs. (1) and (9) it is seen that no mathematical relationship between the two models can be determined. to that end, an iterative approach balancing the area above and below the mc model, see fig. 2, was developed in [1] to get c = σci[(1 + 2a)s + (1 − a) mbσ1n](s + mbσ1n)a−1 a $ 1 + (6amb(s + mbσ1n)a−1)/a , (11) a = (1 + a)(2 + a), sin ϕ = % 6amb(s + mbσ1n)a−1 2(1 + a)(2 + a) + 6amb(s + mbσ1n)a−1 & , (12) for the range of stresses of σt > σ1 > σ1,max, where the bi-axial tensile stress σt = σcis/mb is found by setting σ1 = σ3 = σt in eq. (1), and σ1n = |σ1,max|/σci depends on the upper limit of the confining stress σ1,max seen in fig. 2. it is this parameter we accord our attention to in the next section. figure 2. nonlinear hoek-brown criterion and equivalent mohr-coulomb model in σ1-σ3 plane. 4. simple oedometer test to investigate the influence of the choice of the upper stress limit σ1,max we consider a simple oedometric test. assuming elasticity we get σ1 = σ2 = ν 1 − ν σz , σ3 = σz , (13) 70 vol. 30/2021 hoek-brown and mohr-coulomb failure criteria parameter symbol value unit poisson’s ratio ν 0.3 [−] uniaxial compressive strength σci 20.0 [mpa] hoek-brown constant mi 8.0 [−] geological strength index gsi 30.0 [−] disturbance factor d 0.0 [−] reduced hoek-brown constant mb 0.656680 [−] parameter s 0.000419 [−] parameter a 0.522340 [−] modulus of elasticity em 1414.20 [mpa] table 1. input and calculated parameters of rock mass. figure 3. geometry, boundary and loading conditions of fem model of oedometer. where σ3 < 0 is the applied vertical compressive stress. as mentioned in the previous section, arriving at a comparable response predicted by the mc and hb models using the equivalent shear strength parameters strongly depends on the adopted range of stresses bounded by σt and σ1,max. while σt depends on the hb model parameters, the choice of σ1,max is governed by specific loading conditions. thus an incorrect choice of this parameter may lead to erroneous predictions of the mc model. to see this, we begin with a value σ1n = 0.25 recommended in [4] for triaxial simulations. assuming the values of the hb model provided in table 1 we get from eqs. (11) and (12) c = 649.0 kpa, ϕ = 22.8◦. substituting eq. (13) into the hoek-brown yield function (7) ν 1 − ν σz − σz − σci ! mb ν 1 − ν −σz σci + s "a = 0, (14) gives the limit value of the vertical stress for the first plastic yielding as σz = −16.165 mpa. moving beyond elasticity calls for numerical simulations. to that end, we adopted the geo5 fem software and set up the most simple geometrical model of an oedometer under plane strain conditions consisting of only two constant strain 3-node triangular elements. boundary and loading conditions of the model are evident in fig. 3. the model was loaded by vertical compression such as to pass the elasticity limit point. (a) elastic stress path (b) figure 4. case 1 σ1n = 0.25: a) evolution of stresses based on hb and mc models, b) intersections of elastic stress path with yield surfaces. this simulation clearly demonstrates the influence of the chosen σ1n ratio on the degree of equivalence. unlike the triaxial test2, where the stress state is known, the evolution of principal stresses in oedometer is controlled by the yield function, i.e. by the strength parameters of the rock mass. therefore, the analytical determination of the limit stress value σ1max is not trivial. the consequence of an incorrect choice of σ1n = 0.25 appears in fig. 4(a) plotting the evolution of stresses in terms of invariant stress measures, where σm is the mean stress and j is the square root of the second invariant of the deviatoric stress. clearly, the hb model predicts the deviation from the elastic 2see for example [7] for more details on the hb model verification using triaxial simulations. 71 t. poklopová, v. pavelcová, m. šejnoha acta polytechnica ctu proceedings (a) elastic stress path (b) figure 5. case 2 σ1n = 0.79: a) evolution of stresses based on hb and mc models, b) intersections of elastic stress path with yield surfaces. response precisely at the elastic limit point for the applied load σz = −16.165 mpa, while the mc model still suggests an elastic behavior. the inadequacy of the mc model setting is better seen in fig. 4(b) plotting the elastic stress-path with the onset of yielding identified by the plus sign for both yield surfaces. evidently, the assumed range of stresses for which the equivalence holds, recall fig. 2, is outside the generated stress state controlled by the mc model. the new estimate of σ1n ratio was obtained numerically adopting the hb model to get the maximum lateral principle stress σ1,max = −15.79 mpa for the maximum applied load σz = σ3 = 30 mpa. this gives σ1n = 0.79 and consequently new estimates of strength parameters c, ϕ c = 1345.5 kpa, ϕ = 15.6◦. the corresponding evolution of stresses is shown in fig. 5(a). a reasonable choice of σ1n ratio is further supported in figs. 5(b) and 6(a) comparing the yield surfaces for individual settings. the actual stress path corresponding to both models appears in 6(b) suggesting a close match for the expected stress range. it is worth mentioning that the stress state generated either in oedometer or triaxial apparatus is constant within the sample. however, in real applications a non-homogeneous stress state arises. one thus has to be extremely careful when specifying the expected range of stresses if adopting just one pair elastic stress path (a) (numerical solution) (b) figure 6. a) influence of estimated stress scale, b) evolution of stresses based on hb and mc models. of equivalent shear strength parameters, essentially independent of the actual stress state. nevertheless, for some specific applications simple relations for the determination σ1,max have been proposed in [1]. one such a relationship for the case of slope stability analysis is examined in the next section. 5. slope stability analysis the slope stability analysis typically aims at providing the factor of safety. for the family of elastic perfectly plastic mc models the factor of safety is defined as the ratio of the current shear strength parameter, e.g. cohesion, to the minimum one for which the convergence of the nonlinear problem still exists. this approach is thus not directly applicable to the hb model. however, the authors in [6] introduced an elegant way to overcome this obstacle. so the analysis presented next proceeds in their footsteps. 5.1. hoek-brown model in stability analysis the approach outlined in [6] builds upon a gradual reduction of the strength of a rock mass by modifying the yield function as f (σ1, σ3) = σ1 − σ3 − σci η ! mb −σ1 σci + s "a , (15) 72 vol. 30/2021 hoek-brown and mohr-coulomb failure criteria parameter symbol value unit unit weight γ 25.0 ' kn/m3 ( poisson’s ratio ν 0.3 [−] uniaxial compressive strength σci 30.0 [mpa] hoek-brown constant mi 2.0 [−] geological strength index gsi 5.0 [−] disturbance factor d 0.0 [−] reduced hoek-brown constant mb 0.067225 [−] parameter s 0.000026 [−] parameter a 0.019210 [−] modulus of elasticity em 410.73 [mpa] table 2. input and calculated parameters of rock mass. where η is the reduction factor. a different reduction factor is introduced to reduce the shear strength parameters ϕ, c as tan ϕd = tan ϕc γϕ , cd = cc γc , (16) where ϕc and cc are the actual strength parameters of the subsoil material and γϕ and γc are their corresponding reduction coefficients. a typical setting in the stability analysis assumes γϕ = γc = γ. the objective now is to determine the relationship between the two reduction factors η and γ. this is achieved by constructing a tangent to the hoek-brown yield function at the current state of stress sitting on the yield function and comparing this slope to the slope of the mc yield surface to get η = 1 2 ) ** +γ , 2 − f̃ ′ .///01 + , 1 γ2 − 1 1 −f̃ ′ 22 1 2 − f̃ ′ 22 + f̃ ′ 3 44 5 , (17) where f̃ ′ = ∂f̃ (σ1) ∂σ1 = −mba ! mb −σ1 σci + s "a−1 . (18) notice that η is a function of the current stress state and must be evaluated for every element of the mesh in every loading step separately. for complete derivation of eq. (17) we refer the interested reader to [6]. 5.2. estimating σ1,max in stability analysis the application of the mc model requires setting the value of σ1,max to estimate the shear strength parameters. by performing a large parametric study on a variety of slope geometries and rock mass properties using bishop’s limit state analysis, hoek et al. [1] proposed the following relationship for σ1,max |σ1max| = 0.72σcm ! σcm γh "−0.91 , (19) where h is the height of the slope, γ is the unit weight of the rock mass and σcm is the rock mass strength defined by eq. (10) σcm = 2c cos ϕ 1 − sin ϕ , (20) the rock mass strength expressed using parameters of the hoek-brown model is then given by substituting eqs. (11) and (12) into eq. (20). it holds σcm = σci (mb + 4s − a(mb − 8s))(mb/4 + s)a−1 2(1 + a)(2 + a) . (21) 5.3. practical use of the methods to test reliability of eq. (19), we performed the slope stability analysis of a homogeneous rock slope employing both constitutive models. the geometry of the computational model of a slope with the height of h = 10 m and inclination of 35.5◦ is shown in fig. 7 together with the finite element mesh. the domain was discretized using either linear 3-node (t3) or quadratic 6-node (t6) triangular elements. an average element edge size was about 0.5 m. the material data listed in the upper part of table 2 were taken from [6]. the model parameters presented in the lower part of this table were determined from eqs. (2) (6). figure 7. geometry of fem model, boundary conditions and finite element mesh. the analysis was carried out using standard shear strength parameters reduction approach implemented in geo5 fem. adopting 73 t. poklopová, v. pavelcová, m. šejnoha acta polytechnica ctu proceedings fem tool geo5 fem plaxis 2019 material model hb hb mc mc hb element type t3 t6 t3 t6 t15 safety factor 1.44 1.37 1.42 1.33 1.35 table 3. results of slope stability analysis. (a) (b) (c) (d) figure 8. potential slip surfaces represented by the localized equivalent deviatoric plastic strain: a) hb criterion, t3 elements, b) hb criterion, t6 elements, c) equivalent mc criterion, t3 elements, d) equivalent mc criterion, t6 elements. eqs. (19), (11) and (12) yields |σ1max| = 189 kpa, c = 20 kpa, ϕ = 21◦. in the case of hb model the parameter η was calculated for the prescribed reduction factor γ using eq. (17). the results are summarized in table 3 for both types of approximations. the result acquired by employing the plaxis 2d [8] software is also presented for the sake of comparison. note that it was derived for the 15-node triangular element and a slightly different finite element mesh. as expected, the t3 element delivers with the same mesh refinement the most optimistic predictions. the results provided by the t6 element, the default setting in geo5 fem, are comparable to that of plaxis 2d. also, comparing the hb and mc predictions in terms of both the factor of safety and graphical representation of the slip surfaces in fig. 8 supports eq. (19) in providing a suitable estimate of σ1,max. it should, however, be pointed out that we considered a particularly simple case of a homogeneous slope. thus if possible, the use of the hb model for the slope stability analysis as proposed in [6] is preferable. 6. conclusions the principal objective of this contribution was to highlight the importance of a correct choice of the stress ranges in the definition of equivalent shear strength parameters when replacing the hoek-brown failure model with the mohr-coulomb model. this was demonstrated by running a numerical analysis of a simple oedometer test for two different estimates of σ1n ratio. while the first choice assuming simply the value recommended for triaxial simulations resulted in vastly different predictions when compared to a direct analysis with the hb model, the second choice based on a numerical estimate of the value of σ1,max led to comparable predictions. next, the slope stability analysis was performed taking advantage of the procedure proposed in [6] so that standard approach based on a gradual reduction of strength parameters was possible to apply to both models. although for the simple case of a homogeneous slope the estimate of σ1,max proposed by hoek et al. [1] delivers satisfactory results with the mc model, the approach based on a direct application of the hb model is preferable, particularly in the case of more complex subsoil conditions. acknowledgements the support provided by the sgs project no. sgs20/038/ohk1/1t/11 is gratefully acknowledged. references [1] e. hoek, c. carranza-torres, b. corkum. hoekbrown failure criterion 2002 edition. proceedings of the 5th north american symposium narms-tac 2002. [2] fine-ltd. geo5 fem, user manual. http://www.fine.cz/, 2020. 74 vol. 30/2021 hoek-brown and mohr-coulomb failure criteria [3] e. hoek, e. t. brown. underground excavations in rock. london: institution of mining and metallurgy, 1980. [4] e. hoek, e. t. brown. practical estimates of rock mass strength. international journal of rock mechanics and mining sciences 34.8:1165–1186, 1997. [5] e. hoek, e. t. brown. the hoek-brown failure criterion and gsi 2018 edition. journal of rock mechanics and geotechnical engineering 2018. [6] t. benz, r. schwab, r. kauther, p. a. vermeer. a hoek-brown criterion with instrinsic material strength factorization. international journal of rock mechanics and mining sciences 45:210–222, 2008. [7] t. poklopová. prediction of the response of geotechnical structure using hoek-brown model implementation, verification. master’s thesis, czech technical university in prague, faculty of civil engineering, 2020. in czech. [8] p. ltd. plaxis manual. http://www.plaxis.nl/. 75 doi:10.14311/app.2021.30.0041 acta polytechnica ctu proceedings 30:41–47, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app evaluation of impact of elevated temperatures on young’s modulus of glt beams lucie kucíkováa, ∗, michal šejnohaa, tomáš jandaa, pavel padevěta, guido marsegliab, c a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b university of seville, high technical school of architecture, av. de la reina mercedes, 2, 41012 seville, spain c university of seville, instituto de matemáticas de la universidad de sevilla, (imus), avenida reina mercedes 41012 seville, spain ∗ corresponding author: lucie.kucikova@fsv.cvut.cz abstract. the influence of elevated temperatures on mechanical behavior of glued laminated timber beams is examined on the basis of tensile tests. dog bone samples prepared from beams exposed to fire of variable duration were categorized with respect to the type and position of the failure crack, type and number of discontinuities such as knots, and the level of browning. the acquired experimental results suggest that the wood variability and the effect of growth discontinuities are probably more significant than the effect of elevated temperatures. to support this conclusion, further study is currently under way, exploring samples from the second series of the fire tests. keywords: fire test, glt beam, tensile test, zero strength layer. 1. introduction glued laminated timber (glt) is a promising renewable material enabling wood to be used for more complex constructions. in this regard, the wood combustibility plays an important role which must be taken into account in every design. when fire occurs in the building, the key task becomes a residual load bearing capacity of wooden parts. reduced dimensions of the member cross-section can easily be measured in situ after removing the charcoal. nevertheless, a certain layer is affected by an elevated temperature and is assumed to have zero strength and stiffness. part 1–2 of eurocode 5 defines this layer as d0 = 7 mm for the fire duration longer than 20 min and unprotected surface [1]. some methods to evaluate this layer are presented in [2] depending on the type of load. the authors provide the zero-strength layer for bending in the range of 9.5 to 20.1 mm. in [3] the thickness of 15 mm is suggested to be more appropriate for bending. from that, at least 7 mm below the char-line, i.e. the base of the char layer, the material is expected to have lower load bearing capacity than in the remaining part of the cross-section, which in turn is considered unaffected. the elevated temperatures are assumed to influence the wood in general. the heat treatment affects various material properties depending on the temperature to which the member is exposed. several authors presented a rapid decrease of the modulus of elasticity [4, 5] and compression strength [6] with temperatures exceeding 220◦c and 150◦c, respectively. the present article should contribute to this research effort concentrating on young’s modulus of wood material forming the residual cross-section of glt beams exposed to fire. the article is divided into six sections. following the introductory part 1 we describe in section 2 large scale fire tests. these are supplemented by pilodyn measurements briefly outlined in section 3. the main part of this paper is concerned with small scale tensile tests discussed in section 4 categorized and subsequently evaluated on the grounds of the observed type of failure. the achieved results are further elaborated in section 5. section 6 finally reviews the essential findings. 2. fire test a possible way to examine the effect of elevated temperature on mechanical properties of glt are quick measurements of pilodyn indentation and small scale tensile tests. while the former tests can be carried out directly on the construction site on extinguished beams, the tensile tests performed in laboratory require production of samples identified with a specific position relative to the charred surface. these mechanical tests were preceded by large fire tests described in detail in [7] so only a brief summary is presented here for the sake of clarity. in particular, eight glt beams, with dimensions of 0.10×0.32×2.38 m and containing eight layers of lamellae, were exposed to fire for various duration of 20 min (samples 7 and 8), 30 min (samples 5 and 6), 40 min (samples 3 and 4) and 60 min (samples 1 and 2) and two types of loading curve. the first loading curve was set according to čsn en 1363-1 (samples 1 4) and the second one was set constant at 600◦c 41 http://dx.doi.org/10.14311/app.2021.30.0041 http://ojs.cvut.cz/ojs/index.php/app l. kucíková, m. šejnoha, t. janda et al. acta polytechnica ctu proceedings with the same initial stage as the previous one (samples 5 8). however, the actual furnace temperature exceeded the set temperatures in most cases. the horizontal experimental setup simulated ceiling layout and contained two beams in each test. the beams had to carry their self-weight only, no additional load was applied. a poorly laid stone wool insulation on the top of the beams enabled burning from all four sides. during the test, the temperature distribution within the whole cross-section was continuously measured via thermocouples placed in drilled holes in even-numbered beams. however, the fluctuations occurred on temperature curves of samples 1, 2, 3, and 4, resulting in unknown temperatures at the end of the measurement. consequently, these samples were excluded from the tensile tests, except for sample 4, which was accepted for further examination for the sake of comparison. to prevent the temperature fluctuations for the rest of the tested samples, the hole openings had to be covered by a ceramic blanket (sibral) resulting in three side fire in the middle part of even beams (the beams equipped with thermocouples). the difference between four side and three side fire is evident in figure 4, where in the case of sample 4 the holes were not covered. the pilodyn measurements were carried out along this experimental campaign both prior to and after completing the fire test. with reference to mechanical testing we may invite the interested reader to [7] for the results of three-point bending test performed on burned beams to address the reduction of stiffness and strength of wood exposed to elevated temperature. in this paper, we revisit this issue in the light of the measured young’s moduli only and compare the results provided by pilodyn and tensile tests. 3. pilodyn measurement the indentation by pilodyn 6j device was used to gain estimates of the longitudinal young’s modulus of glt beams both before and after the fire test. the measurement process is based on driving a spike of 2.5 mm in diameter into the wood in transverse direction with energy of 6 j. the only output is the penetration depth d [mm] with an accuracy of 0.5 mm. the value of young’s modulus e [gpa] then follows directly from an empirical equation with only one variable d [mm] as e = a − b d, (1) where parameters a, b are determined experimentally. two such expressions presented in [8] in the form e1 = 19.367 − 0.5641 d, (2) e2 = 20.15 − 0.766 d, (3) are examined in this study. the first equation (2) is generally adopted for wood, whereas eq. (3) was developed for a particular application of glt beams in bending. mind that both equations were proposed for wood not exposed to elevated temperatures. the indentation measurement was performed on all samples before the fire test in the defined grid of 5×8 points per beam located horizontally 0.2 m from the left side with 0.5 m gaps in between and vertically to the middle of each lamella (40 points/beam in total). after the fire test, beams 4, 6, and 8, cleaned from charcoal, were indented in cross pattern (70 points/beam in total). 3.1. results the mean values of measured depth acquired before and after the fire test together with the resulting young’s moduli computed using eqs. (2) and (3) are summarized in table 1. from that, eq. (2) appears more appropriate if no connection to the type of loading and structural application is given. apart from that we noticed that some of the depths measured after the test were too big and the resultant moduli provided by both equations became even negative. therefore, an application of these equations to wood exposed to fire is questionable, as these were originally derived for wood without thermal degradation. n. d [mm] e1 [gpa] e2 [gpa] b a b a b a 4 14.03 17.57 11.46 9.45 9.41 6.69 5 13.85 11.55 9.54 6 13.18 19.36 11.93 8.44 10.06 5.32 7 13.43 11.79 9.87 8 14.16 19.01 11.38 8.64 9.30 5.59 b before fire test; a after fire test table 1. pilodyn 6j indentation measurement. the mean indentation depth increased after the fire test suggesting reduction of young’s modulus. note that the resulting depth could be affected by unevenness of the burnt surface. nevertheless, this reduction was observed when testing different group of samples, where the surface at the location of indents was ground off to eliminate this effect. besides the identified decrease in values of the longitudinal young’s moduli, no correlation with the fire test duration was identified. 4. tensile test small dog bone samples were produced from about 25 cm long segments cut from five glt beams sustaining the fire load (samples 4, 5, 6, 7, 8) and two beams without thermal degradation (control group c). an example of dog bone sample having thickness of about 5 mm is depicted in figure 1(a). the samples were extracted across the width of the beam, with the maximum number of 8 samples depending on residual dimensions. disregarding the outer most lamellae the specimens were produced from two 42 vol. 30/2021 impact of elevated temperatures on glt beams a) dimensions b) failure crack location figure 1. dog bone sample. central lamellae (samples denoted as s) and four other lamellae (samples denoted as k) adjacent in pairs to the central ones. the specimens width was aligned with the lamella height. the wood fibers were oriented parallel to the specimen longitudinal axis. altogether, 406 samples were prepared and tested in the displacement control regime using the mts alliance 30kn electromechanical testing machine equipped with 30kn load cell. a standard extensometer with length of 100 mm was used to monitor the strain of the middle part. the modulus of elasticity is defined as the initial slope of the stress-strain curve. the key output of the tensile test is the load displacement curve. these curves were transformed into stress-strain curves by dividing the load by the initial area measured in the specimen mid section to get the stress and by dividing the extensometer expansion by its initial length (100 mm) to get the strain. all stress-strain curves were examined separately while eliminating those indicating measurement error, e.g. curves with large fluctuations. 4.1. data evaluation during careful visual inspection of all damaged samples, different types of failure modes were identified. comparing the characteristics of individual failures, the specimens were categorized into six groups according to the type of failure crack displayed in figure 2 in several illustrative examples. when the clear sample split into parts with irregular edges with splinters, the failure crack was categorized as ’a along fibers’. if the crack ran directly perpendicular to the fibers, it was denoted as ’b across fibers’. natural wood usually contains a certain amount of growth imperfections. the most pronounced are knots, which in many cases acted as a crack initiators depending on its size. the cracks propagated either across the knot (’e’) or along the knot (’f’), both with a similar frequency, see table 4. the group ’d discontinuity’ covers other discontinuities occurring in wood such as resin pockets, see figure 2, or change in the fiber direction due to vicinity of the knot. the last type (’c delamination’) figure 2. modes of failure. occurred only in a few cases, as the sample split into two plates or layers resembling delamination. further categorization was based on the location of the crack indicated in figure 1(b). if the crack extended to more than one location, the decisive factors were the majority of crack extension into a given location or the major impact. every sample was also rated considering the level of browning ranging from 0 for clear samples to 1 for black. the sample with failure crack running along (’a’) or across (’b’) the fibers and located in the middle (’d’) was considered as the ”ideal” one. 4.2. results the resulting mean and median values of young’s moduli determined from the loading part of the stress-strain curves are presented in tables 2 6 considering the described categorization. the specimens denoted as c were prepared from beams without thermal degradation, used as a control group. recall that the k and s letters refer to the edge and the middle part of the beam, respectively. the statistical data in table 5 were evaluated from both all samples (no elimination) and samples considering the type of failure ’a’ or ’b’ and position ’c’ or ’d’, thus excluding all samples identified with failure modes ’c’, ’d’, ’e’ and ’f’ and failures occurred in ’a’ or along ’b’ grip area (with elimination). the effect of discontinuities and other inaccuracies is, therefore, reduced and the resulting moduli are higher. in all cases, the mean value of (k) samples is lower than the one corresponding to (s) samples. despite eliminating all non ideal specimens no correlation with the position within the beam width has been found. considering the same elimination as mentioned previously, the mean value (9.01 gpa) and the median (9.30 gpa) of the browned samples (26) are only slightly lower than the not browned spec43 l. kucíková, m. šejnoha, t. janda et al. acta polytechnica ctu proceedings n. 4, 5, 6, 7, 8 c browning 0 >0 0 no elimination mean 9.13 8.15 7.18 median 9.06 8.43 7.02 num 282 43 40 with elimination mean 10.18 9.01 8.53 median 9.96 9.30 7.16 num 120 26 15 table 2. young’s moduli from tensile test in [gpa] depending on the level of browning. pos. num. e [gpa] mean median min max a 20 9.18 8.05 5.70 13.89 b 69 9.86 9.48 6.16 17.08 c 140 9.75 9.43 5.78 16.10 d 177 7.67 7.43 0.34 16.19 table 3. young’s moduli from tensile test according to failure position. imens (120) with the mean value (10.18 gpa) and the median (9.96 gpa). without elimination, the mean value (8.15 gpa) and the median (8.43 gpa) of browned samples (43) are still only slightly lower than the not browned specimens (282) with the mean value (9.13 gpa) and the median (9.06 gpa)1, see also table 2. in addition, the control samples which were not exposed to elevated temperatures (denoted as c) experienced the lowest values of young’s moduli compared to the fire test samples, see also table 5. mind that due to the process of sample preparation the most burnt parts of the beam were cut off, therefore, the parts most affected by thermal degradation were excluded. another examined parameter was the contribution of the position of the failure crack and the failure mode. the position of failure crack according to figure 1(b) seems not to affect the resulting young’s modulus. the values of young’s modulus listed in tables 3 and 6 do not differ significantly from each other, even when comparing different positions with the same failure mode. slightly smaller values of position ’d’ in table 3 are caused by a large number of discontinuities (knots and others). considering the most common failure mode ’a’, the mean moduli for all positions are 10.64 gpa (a), 9.74 gpa (b), 9.84 gpa (c), and 10.36 gpa (d), respectively. unfortunately, the number of data in position ’a’ is too small to draw any conclusion. more pronounced are differences among individual modes of failure, recall figure 2. again, comparing the values in tables 4 and 6, the most significant de1the number of samples is in parentheses. mode num. e [gpa] mean median min max a 146 10.07 9.90 1.45 16.19 b 96 9.55 9.16 4.41 17.08 c 10 9.69 8.02 6.72 14.30 d 21 8.21 7.94 4.30 13.86 e 65 5.43 5.42 0.34 15.13 f 68 7.79 7.89 1.53 13.88 table 4. young’s moduli from tensile test according to failure mode. a) a: along fibers b) b: across fibers c) c: delamination d) d: discontinuity e) e: across knot f) f: along knot figure 3. young’s moduli (e [gpa]) from tensile test for individual failure modes. crease is in the case of failure across the knot (mode ’e’) with the mean value of 5.43 gpa. this could be seen in figure 3, where the measured data of mode ’e’ 3(e) fall below those corresponding to mode ’a’ 3(a) or mode ’b’ 3(b). point out that this category (mode ’e’) still comprises all types and sizes of knots, which may lead to low as well as high values (see min and max values in table 4). differentiating between knots and adopting larger statistical data set would reveal more dependencies. because young’s moduli in the transverse direction are lower than in the longitudinal, this decrease is reasonable due to change of growth orientation. however, the values still occur in a similar interval (min, max) as in other failure modes. slightly lower values are observed for the failure along a knot (mode ’f’), where the mean young’s modulus is 7.79 gpa. this category comprises the effect of change in fiber orientation along the knot and an imperfect connection between the knot and the surrounding material. considering the failure along fibers (mode ’a’) and across fibers (mode ’b’) re44 vol. 30/2021 impact of elevated temperatures on glt beams no elimination with elimination time k s k s [min] num mean median num mean median num mean median num mean median c 0 21 5.19 5.89 19 9.38 9.08 10 6.91 6.98 5 11.76 12.86 4 42 28 7.18 8.14 24 9.41 9.70 10 7.83 8.69 14 9.34 9.70 5 31 26 7.57 7.65 29 10.55 10.58 10 9.58 9.34 15 11.91 12.50 6 31 29 8.43 8.04 26 8.46 8.52 17 8.91 8.25 4 9.77 10.21 7 23 37 9.87 10.20 37 10.77 10.48 17 10.59 11.00 15 11.96 12.76 8 23 21 8.80 9.17 25 9.31 10.32 6 9.87 10.21 12 10.73 11.07 table 5. young’s moduli from tensile tests in [gpa] with or without elimination. pos. mod. n. e [gpa] mean median min max a a 4 10.64 11.49 5.70 13.89 a b 3 10.02 8.20 8.09 13.79 a c 1 11.73 11.73 11.73 11.73 a d 2 a e 1 a f 9 7.82 7.75 7.00 9.31 b a 11 9.74 9.84 7.29 12.16 b b 50 9.87 9.42 6.16 17.08 b c 0 b d 0 b e 3 10.84 10.84 10.48 11.21 b f 5 9.48 9.60 9.08 9.76 c a 72 9.84 9.47 5.80 15.85 c b 26 10.25 10.07 6.54 16.10 c c 4 8.22 7.76 6.85 10.50 c d 6 8.82 9.48 6.03 11.27 c e 14 8.69 8.20 5.78 15.13 c f 18 9.81 8.86 7.34 13.88 d a 59 10.36 11.12 1.45 16.19 d b 17 7.52 7.92 4.41 9.73 d c 5 10.65 10.80 6.72 14.30 d d 13 7.93 7.52 4.30 13.86 d e 47 4.55 4.40 0.34 14.13 d f 36 6.77 7.08 1.53 10.53 table 6. young’s moduli from tensile test for combining positions of failure cracks and corresponding failure modes. sulted in similar mean values, namely 10.07 gpa and 9.55 gpa. the remaining modes (’c’ and ’d’) are represented by small data sets, so further conclusion could not be drawn. 5. discussion the results obtained from small-scale tensile tests suggest that the wood variability and the effect of growth discontinuities are probably more significant than the effect of elevated temperature. this is evident when comparing the values of young’s modulus obtained from all samples, recall table 2, and those pertinent to samples free of defects summarized in tables 4 and 6 also identifying the effect of various failure modes. however, we should still keep in mind that producing the specimens requires rubbing down the surface layer. this is because the outline of the beam’s cross-section is rather wavy, see figure 4, and the specimen surface needs to be flat. thus the parts of wood, most affected by elevated temperatures, were essentially removed. nevertheless, a certain amount of browning was observed on 43 specimens (table 2, which indicates temperatures of at least 221◦c (for a light brown) [5]. only four samples evinced dark brown color evaluated as 0.8 (level of browning). not a single black sample with browning level 1 was found. a) sample 4 b) sample 6 c) sample 8 (42.15 min) (31.40 min) (22.90 min) figure 4. temperature distribution [◦c] within cross-section at the end of the fire test (the last measured value before fluctuations occurred is presented in pictures in parentheses). figure 4 shows residual cross-sections in the midspan of the beams with temperatures measured by thermocouples at the end of the fire test. the numbers in parentheses denote the last measured values before fluctuations occurred on temperature curves. the temperatures close to the cross-section outline are relatively low, which also suggests a rather smaller thickness of the zero-strength layer, reminding a rapid decrease of the modulus of elasticity above 220◦c 45 l. kucíková, m. šejnoha, t. janda et al. acta polytechnica ctu proceedings [4, 5]. however, the values on the edges of the crosssection are below a widely accepted temperatures of 300◦c [1] or 288◦c (550◦f) [9, 10], which defines the base of the char layer (i.e. pyrolysis temperature). this could mean either position error within the picture or that the charring starts at lower temperatures, where e.g. lange et al. [3] defined the char front already at 270◦c. nonetheless, the results of the pilodyn indentation in table 1 shows a certain decrease of the longitudinal moduli. even if the presented equations (2) and (3) were not applicable to wood loaded by fire, the indentation depths after the fire test are still higher than that obtained before the test. therefore, an assumption of the softening of surface layers of charred beams is reasonable. 6. conclusions the tensile test was performed on small dog bone samples, produced from five beams sustaining the fire load and two beams without thermal degradation, to obtain the longitudinal young’s moduli. the samples were extracted across the beam width from two locations (k) and (s) samples. the specimens were also categorized into groups according to a defined failure mode, figure 2, and the failure position, figure 1(b), and they were also rated considering the level of browning (0–without browning, 1–black). the resulting young’s moduli are stated in table 5 considering all samples as well samples upon elimination of those with the failure crack located in or along the knot or other discontinuity and those occurring in or along the grip area. despite elimination, no correlation with position within the beam width has been found. the browned samples showed only slightly lower values than the not browned specimens, recall table 2. in addition, the control samples not exposed to elevated temperatures evinced the lowest values of young’s moduli. on the contrary, larger decrease of the moduli was observed by pilodyn 6j indentation. the differences between values measured before and after the fire test, see table 1, lead to an assumption that at least the properties of the surface layers are reduced. but keep in mind that eqs. (2) and (3) were originally derived for members not exposed to elevated temperature and their applicability could be considered questionable. nevertheless, the increased indentation depth still shows some level of softening of the surface. the largest reduction of the measured young’s moduli was observed for failure modes where the failure crack propagated across the knot (mode ’e’), see tables 4 and 6, and figure 3. therefore, the presence of knots affects the resulting material properties more than the elevated temperatures. less pronounced effect was also observed for modes ’f’ (along knot) and ’d’ (discontinuity). on the other hand, no correlation with the position within the dog bone sample was found as evident from tables 3 and 6. the results obtained from the small-scale tensile tests suggest that the wood variability and the effect of growth discontinuities are probably more significant than the effect of elevated temperature. to support this conclusion, further study is currently under way exploring samples from the second series of the fire tests. list of symbols e young’s modulus [gpa] d pilodyn 6j indentation depth [mm] k samples from edges (cross-section height point of view) s samples from middle (cross-section height point of view) c control samples a − d failure position (figure 1) a − f modes of failure (figure 2) acknowledgements the financial support provided by the czech technical university in prague within sgs project with the application registered under the no. sgs20/155/ohk1/3t/11 and by the gačr grant no. 18-05791s is gratefully acknowledged. references [1] european committee for standardization. eurocode 5: design of timber structures part 1-2: general structural fire design, 2004. [2] j. schmid, a. just, m. klippel, m. fragiacomo. the reduced cross-section method for evaluation of the fire resistance of timber members: discussion and determination of the zero-strength layer. fire technology 51(6):1285–1309, 2015. doi:10.1007/s10694-014-0421-6. [3] d. lange, l. boström, j. schmid, j. albrektsson. the influence of parametric fire scenarios on structural timber performance and reliability. tech. rep., sp technical research institute of sweden, 2014. [4] b. esteves, h. pereira. wood modification by heat treatment: a review. bioresources 4(1):370–404, 2008. [5] d. kačíková, f. kačík, i. čabalová, j. ďurkovič. effects of thermal treatment on chemical, mechanical and colour traits in norway spruce wood. bioresource technology 144:669–674, 2013. doi:10.1016/j.biortech.2013.06.110. [6] s. yildiz, e. d. gezer, u. c. yildiz. mechanical and chemical behavior of spruce wood modified by heat. building and environment 41(12):1762–1766, 2006. doi:10.1016/j.buildenv.2005.07.017. [7] l. kucíková, t. janda, m. šejnoha, j. sýkora. experimental investigation of fire resistance of glt beams. international journal of computational methods and experimental measurements 8(2):99–110, 2020. doi:10.2495/cmem-v8-n2-99-110. 46 https://doi.org/10.1007/s10694-014-0421-6 https://doi.org/10.1016/j.biortech.2013.06.110 https://doi.org/10.1016/j.buildenv.2005.07.017 https://doi.org/10.2495/cmem-v8-n2-99-110 vol. 30/2021 impact of elevated temperatures on glt beams [8] m. šejnoha, t. janda, l. melzerová, et al. modeling glulams in linear range with parameters updated using bayesian inference. engineering structures 138:293– 307, 2017. doi:10.1016/j.engstruct.2017.02.021. [9] american wood council. calculating the fire resistance of wood members and assemblies, technical report no. 10, 2020. [10] r. h. white, e. v. nordheim. charring rate of wood for astm e 119 exposure. fire technology 28(1):5–30, 1992. doi:10.1007/bf01858049. 47 https://doi.org/10.1016/j.engstruct.2017.02.021 https://doi.org/10.1007/bf01858049 acta polytechnica doi:10.14311/app.2020.27.0037 acta polytechnica 27(0):37–41, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app dynamic impact wear and impact resistance of w-b-c coatings josef daniela, ∗, jan grossman a, vilma buršíková b, lukáš zábranský b, pavel součekb, saeed mirzaeib, petr vašinab a institute of scientific instruments of the czech academy of sciences, královopolská 147, brno 612 64, czech republic b masaryk university, faculty of science, department of physical electronics, kotlářská 2, brno 611 37, czech republic ∗ corresponding author: jdaniel@isibrno.cz abstract. coated components used in industry are often exposed to repetitive dynamic impact load. the dynamic impact test is a suitable method for the study of thin protective coatings under such conditions. aim of this paper is to describe the method of dynamic impact testing and the novel concepts of evaluation of the impact test results, such as the impact resistance and the impact deformation rate. all of the presented results were obtained by testing two w-b-c coatings with different c/w ratio. different impact test results are discussed with respect to the coatings microstructure, the chemical and phase composition, and the mechanical properties. it is shown that coating adhesion to the hss substrate played a crucial role in the coatings’ impact lifetime. keywords: boron carbide, fracture resistance, impact resistance, impact wear, w-b-c. 1. introduction industrially used components coated with protective coatings are often exposed to the external load. in some cases, this load may take the form of repeated dynamic impact. for the investigation of the lifetime of the coated components under this form of load, knotek et al. developed the method of dynamic impact test in the early 90s [1]. a tungsten carbide ball repeatedly impacted the surface of the specimen. the load force of the ball was well-defined, and the impact frequency was set to a constant value. coatings’ lifetime was evaluated as a maximum number of impacts before coating failure. knotek also showed that dynamic impact test makes a study of dynamic wear and fatigue behaviour of the whole coating/substrate system possible [1]. in the past decades, the dynamic impact testing was investigated by different researchers. bouzakis et. al. developed fem simulation of the elastoplastic deformations of material under dynamic impact load [2] or the inclined dynamic impact test testing of the coated specimens under various angles of inclination [3]. engel et. al. proposed a three-stage model of the dependence of the impact crater volume on the total number of impacts based on the coating deformation [4]. voevodin et. al. investigated the impact lifetime of various ti-based coatings and discussed it with respect to their chemical composition and hardness [5]. heinke et. al. used the dynamic impact test as a supplementary method to the evaluation of coatings adhesion [6]. sobota et. al. investigated the behaviour of dlc coatings under dynamic load in various air humidity [7]. in the past years, the dynamic impact testing also expanded to the field of the thick hvof sprayed coatings [8]. the aim of this paper is to present some crucial results of dynamic impact testing, important for the investigation of the behavior of the coating under dynamic impact load and for development of the new types of the protective coatings. novel parameters of impact resistance and impact deformation rate, suitable for better understanding of the behaviour of the coating under repeated impact load, were discussed. presented results were obtained from two w-b-c coatings with different chemical composition, structure, and mechanical properties. 2. experimental 2.1. dynamic impact tester the coatings were tested by a dynamic impact tester developed at the institute of scientific instruments of the czech academy of sciences in brno. the impact hammer with a lift set at a constant value of 2.4 mm was electromagnetically driven. attached to the impact hammer was a ball-shaped cemented tungsten carbide impact indenter with a diameter of 5 mm. the indenter ball was rotated to an unworn contact side before every test. the frequency of impact testing was set to 8 hz, and the duration of the contact of the ball indenter with the surface was 1.5 ms. the impact testing was carried out with impact loads of 200 n, 400 n and 600 n. due to its construction the impact tester is uniquely suitable also for the investigation of the behaviour of the samples under a low number of impacts (≤ 100). therefore, the number of impacts was in the range from 1 to 500 000. every test was repeated three times. the depth, 37 http://dx.doi.org/10.14311/app.2020.27.0037 http://ojs.cvut.cz/ojs/index.php/app j. daniel, j. grossman, v. buršíková et al. acta polytechnica c/w w (at. %) c (at. %) b (at. %) h (gpa) e (gpa) h/e h3/e2 (gpa) thick.(µm) 0.11 63 7 30 28.6 ± 1.5 500 ± 12 0.057 0.094 1.7 0.8 40 32 28 23.3 ± 1.0 400 ± 12 0.058 0.079 1.3 table 1. chemical composition and mechanical properties of the w-b-c coatings. figure 1. loading curves of the w-b-c coatings. dotted lines were added as a visual guide. the radius and the surface morphology of the impact craters were measured by talystep profilometer and lext ols 3100 confocal microscope. 2.2. w-b-c coatings preparation two w-b-c coatings with different c/w ratios were deposited on hss substrate by magnetron cosputtering of w, c and b4c targets in ar atmosphere. for the detailed description of the coatings’ deposition, see ref. [9]. 3. results 3.1. w-b-c coatings composition, structure and mechanical properties the chemical composition and the mechanical properties of the presented w-b-c coatings are summarized in table 1. both coatings exhibited an amorphous microstructure. the coating with a higher amount of carbon (c/w = 0.80) exhibited a higher amount of c-c bonds and b-c bonds and a lower amount of w-b bonds. the detailed description of the chemical composition and the microstructure analyses, the evaluation of the mechanical properties and the discussion of the relationship between the microstructure and the mechanical properties can be found in ref. [9]. 3.2. results of impact testing one of the first results of the dynamic impact test is a loading curve the dependence of the impact crater volume on the total number of impacts. the loading curves of the studied w-b-c coatings are depicted in figure 1. the amount of energy supplied from the tester to the tested specimen increased with the increasing number of impacts. this energy caused the deformation of the specimen and thus the formation of the impact crater. as the number of impacts increased further, the supplied energy was dissipated in the form of a continuously increasing internal stress. after reaching some critical number of impacts nc, the tested coating/substrate system was unable to receive more energy from the tester, and the coating failed. the depth and the radius and thus the volume of the impact crater then rapidly increased. these different behaviour modes of the substrate/coating system can be seen on the loading curve as different slopes and thus it is often possible to determine the nc and therefore the lifetime of the coating under dynamic impact load. the impact craters in the w-b-c coatings formed after 10 000 impacts are depicted in figure 2. the coating with the c/w ratio of 0.11 exhibited impact craters with a higher diameter than the coating with the c/w ratio of 0.80. moreover, in case of the load of 400 n, this coating exhibited an almost uncovered substrate indicating the first stage of the coating failure. the total coating failure and the uncovered substrate in case of the load of 600 n for the coating with the c/w ratio of 0.11 corresponded to a rapid increase of the slope of the loading curve of this coating depicted in figure 1. while the coating with the low c/w ratio exhibited pronounced chipping cracks in the case of the loads of 400 n and 600 n, the coating with the higher c/w ratio only exhibited small concentered cracks even in case of the highest load of 600 n. the detailed view on the edges of impact craters created with load of 400 n is depicted in figure 3. another parameter describing the behaviour of a coated specimen under the dynamic load is the impact resistance. the impact resistance is defined as follows. the volume of the impact crater created by n1 impacts is denoted as v1 and the volume of the impact crater created by n2, impacts is denoted as v2. an increment of the impact crater volume ∆v = v2 − v1. this 38 vol. 27/2020 dynamic impact wear and impact resistance of w-b-c coatings figure 2. comparison of impact craters created by 10 000 impacts. figure 3. detail view on the edges of impact craters created with load of 400 n and 10 000 impacts. increment depends on the number of impacts n1 and n2 [4]. assuming that the increment of ∆v is inversely related to the impact resistance of a coating, the impact resistance can be calculated from the volume of impact craters at a given number of impacts. this parameter with the unit of [ n/mm3 ] describes how many impacts (denoted as n ) are required to create a crater with the volume of 1mm3. the impact resistance of the w-b-c coatings is shown in figure 4. at first, the impact resistance increased with the increasing crater volume due to the elastoplastic deformation of coatings. as the number of impacts increased further, the energy from the tester was dissipated in the form of internal stress and not in the form of the crater deformation. the impact resistance thus reached the highest values. then, as the coating failed, the volume of the impact crater and also the increment of the volume ∆v rapidly increased, and the impact resistance correspondingly rapidly decreased. therefore, the critical number of impacts nc can also be evaluated using the impact resistance. the third way of evaluation of the critical number of impacts is from the determination of the deformation rate. assuming that contact time between the impactor ball and the specimen is the same for every impact, it is possible to calculate the deformation rate as the impact crater volume divided by the total contact time. the deformation rate of the w-b-c coatings is depicted in a log-log scale in figure 5. the deformation rate linearly decreased as the number of impacts increased. after reaching the critical number of impacts, the impact crater volume rapidly increased and thus deformation rate curve deviated from linearity. the final critical number of impacts was evaluated from the combination of the previously described methods from the loading curves, the optical investigation of impact craters (uncovered substrate as evidence of the coating failure), the impact resistance and the deformation rates. figure 6 shows the woehler-like dependence of the critical number of impacts on the load. the coating with the c/w ratio of 0.80 exhibited higher lifetime under dynamic impact at all of the applied loads. 4. discussion in general, the behaviour of coatings under impact load is closely related to the coating’s toughness [2, 5]. both w-b-c coatings exhibited an amorphous structure and the same h/e ratio that is related to the fracture resistance [9]. the coating with the c/w ratio of 0.11 exhibited higher hardness and elastic 39 j. daniel, j. grossman, v. buršíková et al. acta polytechnica figure 4. impact resistance of the w-b-c coatings. dotted lines were added as a visual guide. figure 5. deformation rate of the w-b-c coatings. figure 6. woehler-like curves of the w-b-c coatings. dotted lines were added as a visual guide. modulus and therefore, a higher h3/e2 ratio that is related to resistance to plastic deformation and the coating toughness [10]. however, better impact resistance was observed in the case of the coating with the c/w ratio of 0.80. the main reason was different adhesion of the coatings to the hss substrate. as was shown in figure 3, the impact crater on the coating with the c/w ratio of 0.11 created after 10 000 impacts with the load of 400 n exhibited cracking on its edge, while the impact crater on the coating with the c/w ratio of 0.80 created after the same number of impacts at the same load exhibited an uninterrupted coated edge. thus, the adhesion of the coating with the c/w ratio of 0.80 was higher than the adhesion of the coating with the c/w ratio of 0.11. as the h/e ratio was the same for both coatings, the adhesion of the studied coatings was the main parameters leading to better impact resistance. 5. conclusions two w-b-c coatings with 30 at. % of boron and different c/w ratios were deposited on hss substrates and analyzed using the dynamic impact test. the impact loads of 200 n, 400 n and 600 n were used. the critical number of impacts determining coating failure was estimated using the loading curves, optical investigations and two novel parameters – impact resistance and deformation rate. the higher adhesion and of the coating to the hss substrate led to the higher lifetime of the coating under repeated impact load. acknowledgements this work was performed under the support of the projects lo1212 and lo1411 funded by ministry of education, youth, and sports of the czech republic, and project ga19-03899s funded by the czech science foundation. references [1] o. knotek. a new technique for testing the impact load of thin films: the coating impact test. surface and coatings technology 54-55:102–107, 1992. doi:10.1016/0257-8972(92)90147-3. [2] k.-d. bouzakis, n. vidakis, t. leyendecker, et al. determination of the fatigue properties of multilayer pvd coatings on various substrates, based on the impact test and its fem simulation. thin solid films 308309:315–322, 1997. doi:10.1016/s0040-6090(97)00561-0. [3] k.-d. bouzakis, a. asimakopoulos, n. michailidis, et al. the inclined impact test, an efficient method to characterize coatings’ cohesion and adhesion properties. thin solid films 469-470:254 – 262, 2004. proceedings of the 31st international conference on metallurgical coatings and thin films, doi:10.1016/j.tsf.2004.08.093. [4] p. engel. impact wear of multiplated electrical contacts. wear 181-183:730–742, 1995. doi:10.1016/0043-1648(94)07105-5. [5] a. voevodin, r. bantle, a. matthews. dynamic impact wear of ticxny and ti-dlc composite coatings. wear 185(1-2):151–157, 1995. doi:10.1016/0043-1648(95)06603-9. 40 https://doi.org/10.1016/0257-8972(92)90147-3 https://doi.org/10.1016/s0040-6090(97)00561-0 https://doi.org/10.1016/j.tsf.2004.08.093 https://doi.org/10.1016/0043-1648(94)07105-5 https://doi.org/10.1016/0043-1648(95)06603-9 vol. 27/2020 dynamic impact wear and impact resistance of w-b-c coatings [6] w. heinke, a. leyland, a. matthews, et al. evaluation of pvd nitride coatings, using impact, scratch and rockwell-c adhesion tests. thin solid films 270(12):431–438, 1995. doi:10.1016/0040-6090(95)06934-8. [7] j. sobota, j. grossman, v. bursikova, et al. evaluation of hardness, tribological behaviour and impact load of carbon-based hard composite coatings exposed to the influence of humidity. diamond and related materials 20(4):596–599, 2011. doi:10.1016/j.diamond.2011.01.011. [8] k. bobzin, l. zhao, m. öte, et al. impact wear of an hvof-sprayed cr 3 c 2 -nicr coating. international journal of refractory metals and hard materials 70:191–196, 2018. doi:10.1016/j.ijrmhm.2017.10.011. [9] m. alishahi, s. mirzaei, p. souček, et al. evolution of structure and mechanical properties of hard yet fracture resistant w-b-c coatings with varying c/w ratio. surface and coatings technology 340:103–111, 2018. doi:10.1016/j.surfcoat.2018.02.054. [10] j. musil. hard and superhard nanocomposite coatings. surface and coatings technology 125(13):322–330, 2000. doi:10.1016/s0257-8972(99)00586-1. 41 https://doi.org/10.1016/0040-6090(95)06934-8 https://doi.org/10.1016/j.diamond.2011.01.011 https://doi.org/10.1016/j.ijrmhm.2017.10.011 https://doi.org/10.1016/j.surfcoat.2018.02.054 https://doi.org/10.1016/s0257-8972(99)00586-1 acta polytechnica 27(0):37–41, 2020 1 introduction 2 experimental 2.1 dynamic impact tester 2.2 w-b-c coatings preparation 3 results 3.1 w-b-c coatings composition, structure and mechanical properties 3.2 results of impact testing 4 discussion 5 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0329 acta polytechnica ctu proceedings 33:329–336, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague correlation of porosity and properties of recycled fine aggregate concrete with fly ash hangwei lina, ∗, koji takasub, hidehiro koyamadac, hiroki suyamac a the university of kitakyushu, faculty of environmental engineering, department of architecture, 1-1 hibikino, wakamatsu, kitakyushu city, fukuoka, 808-0135 japan b the university of kitakyushu, faculty of environmental engineering, 1-1 hibikino, wakamatsu, kitakyushu city, fukuoka, 808-0135 japan c the university of kitakyushu, department of architecture, faculty of environmental engineering, 1-1 hibikino, wakamatsu, kitakyushu city, fukuoka, 808-0135 japan ∗ corresponding author: 615299569@qq.com abstract. in this paper, we presented the experimental discussion of samples of recycled aggregate concrete with replacement of natural fine aggregate by recycled fine aggregate. three mix kinds were produced and, for each of these three kinds, two levels of water to blinder ratio were used with 0%, 10% and 20% of fa. the result of the tests of drying shrinkage and compressive strength of recycled concrete were used for comparison with tests of mercury intrusion porosimetry (mip), in which the cumulative pore volume and different intervals of pore volume were studied at ages of 28 and 91 days. correlation graph of compressive strength and cumulative pore volume might be predicted with given w/b, days and fa of the concrete or none given and drying shrinkage might be predicted with given fa. keywords: fly ash, drying shrinkage, porosity, recycled aggregate concrete. 1. introduction at present, the decrease in strength and increase in drying shrinkage are major problems for recycled aggregate concrete, and it was considered that managing these will lead to an increase in the use of recycled aggregate concrete (rac) to building structures. many studies have shown that using recycled aggregates in concrete suffered from durability problems. some people reported that, compared with the reference concrete, the shrinkage of recycled aggregate concrete was increased by up to 60% [1]. in past research, the replacement rate for replacing natural aggregate with recycled aggregate (ra) and replacing cement with fly ash (fa) were considered roughly or simply. in this research, i focused on the compressive strength and drying shrinkage of concrete both with fine recycled aggregate and fly ash. then i made correlation graphs between porosity and properties of the concretes. then the correlation process was carried out on the properties obtained from the mercury intrusion porosimetry (mip) tests, and shrinkage and compressive strength. finally carrying out the graphs that gave rise to the correlations of the shrinkage and compressive strength and the porosity obtained in the mercury intrusion porosimetry mip tests. 1.1. experimental outlines table 1 showed the physical properties of the materials used in this experiment. as cement, ordinary portland cement was used, and sea sand and recycled fine aggregate as the fine aggregate (m standard of jis a 5022) and crushed stone as coarse aggregate were used. table 3 showed the mix proportions. in the mix proportions, the unit water amount was 180kg/m3, the unit cement amount was 327kg/m3, the water to binder ratio was 55% and 44%, and the replacement rate of the regenerated fine aggregate was 2 levels of 50, 100%. in this experiment, recycled aggregate was used only for fine aggregate, and natural aggregate crushed stone was used for coarse aggregate in all mix proportions. regarding the compounding symbols, n was the mix proportions using natural aggregate for both the coarse aggregate and the fine aggregate. the replacement rate of the recycled fine aggregate was indicated in this order, m50 was a mix proportions using 50% recycled fine aggregate class m and 50% sea sand. the replacement rate of the fly ash was indicated in this order, n-fa10 was a mix proportions replacing cement with 10% fly ash. the compressive strength test was carried out according to jis a 1108. the sample for the mip rests were extracted from the center of the cylinder (! = 0.05 m × h = 0.1 m) which were dried by putting them in curing room for 28 and 91 days. the cylinder was then cut with a vise to obtain particles from 2 mm to 5 mm in diameter. the samples were dehydrated by submerging them in acetone for 2 h. after this stage, the samples were dried by putting them in an oven at 105 ◦c for 1 h. finally, the samples were put into a vacuum pump for 48 h. the mip tests were done on autopore v which subjected the samples to maximum pressure of 329 https://doi.org/10.14311/app.2022.33.0329 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en h. lin, k. takasu, h. koyamada, h. suyama acta polytechnica ctu proceedings property coarseaggregate fine aggregate fine recycled aggregate contents cement fa density (g/cm3) 2.69 2.59 2.23 loss of ignition (%) 2.40 2.1 fineness modulus 6.9 2.58 4.3 density(g/cm3) 3.16 2.2 water absorption (%) 1.14 1.04 7.38 specific surface area (cm2/g) 3310 3270 table 1. materials. w/b (%): 55 n m50 m100 n-fa10 m50-fa10 m100-fa10 n-fa20 m50-fa20 m100-fa20 fa 0 0 0 45 45 45 90 90 90 s 731 350 0 723 346 0 716 342 0 rfa 0 350 672 0 346 665 0 342 658 w/b (%): 40 n m50 m100 n-fa10 m50-fa10 m100-fa10 n-fa20 m50-fa20 m100-fa20 fa 0 0 0 33 33 33 65 65 65 s 832 398 0 826 396 0 820 392 0 rfa 0 398 765 0 396 759 0 392 754 g 945 w 180 table 2. unit mass of mix proportions. up to 60,000 pisa. 2. experimental results and discussion 2.1. compressive strength 2.1.1. influence of fine ra on the compressive strength figure 1 showed the compressive strength at 28 and 91 days for concrete mixes with various levels of fine ra. for 28 days, it showed that incorporating fine ra in concrete mixes decreased the compressive strength in fa10% w/c 0.40 series and fa 20% w/c0.55 series. but other four series showed that incorporating fine ra in concrete mixes almost had no decreased the compressive strength. for 91 days, it showed that incorporating of fine ra in concrete mixes decreased the compressive strength in w/b-0.40 fa-10% series and w/b-0.55 fa-20% series. and the incorporating of fine ra in concrete mixes made no effect on the compressive strength in w/b-0.40 fa-0% series and w/b-0.55 fa-10% series. this might be explained by two main factors, namely the rca being more porous and weaker than the na [2] and/or due to the increase of the effective w/c ratio in rca concrete mixes to maintain the workability of an equivalent conventional concrete [3]. 2.1.2. influence of fa on the compressive strength figure 2 showed the compressive strength at 28 and 91days for concrete mixes with various levels of fa. for 28 days, it showed that the incorporation of fa in concrete mixes was detrimental in terms of compressive strength, but slightly decreased. for 91 days, it showed that the incorporating of fa in concrete mixes made no effect on increasing the compressive strength on w/b-0.40 ra-100% series, w/b-0.55 ra-100% series and w/b-0.55 ra-0% series. generally, the incorporation of a low volume of fa filled the voids between cement particles, thus decreasing the voids, and slightly increased the fresh density. this was likely to happen because fa particles are finer than cements, so that fa not only can improve the particle size distribution, but also filled the micro pores in the interfacial transition zones around the aggregates [4]. 2.2. drying shrinkage 2.2.1. influence of fine ra on the drying shrinkage figure 3 showed the drying shrinkage at 182 days for concrete mixes with various levels of fine ra. each presented value was the average of three measurements. the results showed that incorporating fine 330 vol. 33/2022 porosity and properties of fly ash concrete figure 1. effect of fine ra on the compressive strength. figure 2. effect of fa on the compressive strength. ra in concrete mixes increased the drying shrinkage of the concretes in w/c-0.40 fa-0% series and w/c-0.40 fa-10% series. on w/c-0.40 fa-0% series, 50% fine ra replacement ratio made more effect on increasing concrete’ durability than 100% fine ra replacement ratio. this was probably the result of the combination of half recycled aggregate and half normal aggregate. the mortar adhered to the recycled aggregate contributed to an increase in the volume of the paste (old + new), thus increasing the drying shrinkage of the resulting concrete [5]. 2.2.2. influence of fa on the drying shrinkage figure 4 showed the drying shrinkage at 182 days for concrete mixes with various levels of fa. the results showed that incorporating fa in concrete mixes decreased the drying shrinkage in w/c-0.40 ra-0% series and w/c-0.40 ra-50% series. both in two series, 10% fa replacement ratio made more effect on decreasing concrete’ durability than 20% fa replacement ratio. this was probably the result of high fa replacement ratio. according to atis et al. [6], the refigure 3. effect of fine ra on the drying shrinkage. figure 4. effect of fa on the drying shrinkage. duction in the drying shrinkage was attributed to the dilution effect of the fly ash particles. on w/c-0.40 ra-100% series, 10% fa replacement ratio would decrease concrete’s durability. this was same as other two series. when 20% fa was replaced, the fa made no effect on decreasing the drying shrinkage of the concrete. it might the that fly ash and recycled aggregates will inhibit each other when they both in high replacement ratio. figure 6 had water to binder ratio 0.40. at 28 days, only 100% fine ra replacement ratio cause the increase of the pore volume when pore diameter was in 0.003 ∼ 0.03 µm in 0% fa replacement ratio series. in 10% fa replacement ratio series, 50% fine ra replacement ratio had the biggest influence on pore volume in 0.003 ∼ 100 µm pore diameter interval. at 91 days, the results showed that the cumulative pore volume increased with the increase of the fine ra replacement ratio when pore diameter was in 0.003 ∼ 100 µm except 0% fa replacement ratio series. 2.2.3. correlation of porosity and properties the pore diameter was divided into many different intervals by its different changed points of the cumulative curve. in figure 7, for example, each line 331 h. lin, k. takasu, h. koyamada, h. suyama acta polytechnica ctu proceedings figure 5. influence of fine ra replacement ratio on cumulative pore volume(w/b0.55). figure 6. influence of fine ra replacement ratio on cumulative pore volume(w/b0.40). has two changed points which were named after point a and b. the compressive strength has relationship with pore diameter interval which left endpoint was after point a. the drying shrinkage has relationship with pore diameter interval which right endpoint was before point b. 2.3. porosity 2.3.1. result of porosity the tests were carried out at 28 and 90 days. figure 5 showed the influence of recycled aggregate replacement ratio on cumulative pore volume with water to binder ratio 0.55. at 28 days, the results showed that the pore volume, when pore diameter was in 0.003 ∼ 100 µm, increased with the increase of the fine ra replacement ratio. at 91 days, the results showed that the pore volume, when pore diameter was in 0.003 ∼ 100 µm, increased with the increase of the fine ra replacement ratio, except 0% fa replacement ratio series. 50% fine ra replacement ratio made more effect on increase of pore volume when pore diameter was in 0.2 ∼ 100 µm in fa 0% replacement ratio series. table ?? showed correlation coefficients (r2) when using different interval division methods (w/b, days and fa) and after the use of different interval division methods, it seemed that w/b, days and fa had no influence on correlations of the compressive strength and pore volume. and the best interval’s endpoint for the pore diameter were 0.05 µm and 100 µm. figure 8 showed the graph that gave rise to the correlations of the compressive strength and the 0.05 ∼ 100 µm pore volume obtained in the mip tests (28 and 91days). y = −3.94x + 54.60, r2 = 0.45 (1) • y: compressive strength (mpa). • x: 0.05 ∼ 100 µm pore volume (×10−2 cc/cc) y = −2.48x + 57.28, r2 = 0.32 (2) • y: compressive strength (mpa) • x: cumulative pore volume (×10−2 cc/cc) the correlation process was carried out on the drying shrinkage of the recycled concrete (91 days). the correlation between porosity and drying shrinkage of the concretes can be divided two degrees: recycled concrete without fly ash and with fly ash replacement. 332 vol. 33/2022 porosity and properties of fly ash concrete figure 7. ’changed points’ of the cumulative curve. figure 8. coefficients of determination (r2) of compressive strength and cumulative pore volume at interval division: (a) a−b µm; (b) a−36 µm; (c) 0.02 − 0.2 µm; (d) 0.02 − 36 µm. 333 h. lin, k. takasu, h. koyamada, h. suyama acta polytechnica ctu proceedings pore volume w/b, days, fa 0.55, 28, 0% 0.55, 91, 0% 0.55, 28, 10% 0.55, 91, 10% 0.55, 28, 20% 0.55, 91, 20% a ∼ b µm r2 − 0.08 0.87 0.07 0.99 0.91 w/b, days, fa 0.40, 28, 0% 0.40, 91, 0% 0.40, 28, 10% 0.40, 91, 10% 0.40, 28, 20% 0.40, 91, 20% r2 0.99 0.58 0.66 0.54 0.01 0.64 w/b, fa 0.55, 0% 0.55, 10% 0.55, 20% 0.40, 0% 0.40, 10% 0.40, 20% r2 0.53 0.31 0.29 0.00 0.49 0.02 a ∼ 2 µm r2 0.10 0.25 0.26 0.00 0.08 0.04 a ∼ 100 µm r2 0.12 0.22 0.15 0.05 0.09 0.03 table 3. correlation coefficients (r2). y = 52.38x + 371.79, r2 = 0.89 (3) • y: drying shrinkage (µ) (with fa) • x: 0.003 ∼ 0.1 µm pore volume (×10−2 cc/cc) y = 48.62x + 257.28, r2 = 0.71 (4) • y: drying shrinkage (µ) (with fa) • x: cumulative pore volume (×10−2 cc/cc) the researchers studied the porosity and pore structure of cement paste. z.w. wu thought pores have four intervals: < 20nm which is named after harmless pore, 20-50nm which is named after lessharm pore, 50-200nm which is named after harmful pore, > 200nm which named after much-harm pore [7]. capillary pores with a diameter greater than 0.011 µmaffect the strength, and gel pores with a diameter of fewer than 0.01 µm affect the durability of mortar [8]. consequently, the correlation between the properties of the rac and the pore volume can be evaluated by dividing various pore intervals. wu’s theory was used for comparison; and the left and right points of the pore diameter intervals were defined as 0.02 and 0.2 µm, respectively. simple linear regression also was performed to evaluate the correlation between the results of mip and the results of compressive strength or drying shrinkage tests under four interval division methods according to the object is strength or dry shrinkage. figure 8 shows the coefficients of determination between the pore volume and compressive strength at the following interval divisions: a−b µm, a−36 µm, 0.02 − 0.2 µm, 0.02 − 36 µm. the pore size range of the mip test was 0.003 − 36 µm. the symbols "all", "fa", "0.55fa" and "0.40fa" shown in figure 8 indicate that the fitted line used the data of the compressive strength and cumulative pore volume of all mixes, mixes with fa, mixes with fa at w/b 0.55 and mixes with fa at w/b 0.40, respectively. for rfa concrete incorporating fa, there was no interval with a high correlation coefficient between the compressive strength and the cumulative pore volume because the correlation coefficients of the four types of intervals ranged between 0 and 0.6 (figure 8). in other words, there was no significant correlation between cumulative pore volume and compressive strength. moreover, simple linear regression cannot use cumulative pore volume in a specific interval to determine or predict the strength of recycled concrete. some researchers use quadratic or exponential regression equations to predict concrete strength. concrete with fa should be considered separately when considering the correlation between compressive strength and cumulative pore volume because the correlation coefficients of the "fa" lines are higher than that of the "all" line for the four intervals. figure 9 shows the coefficients of determination between the cumulative pore volume and drying shrinkage at interval divisions: a−b, a − 36 µm, 0.02 − 0.2 µm, 0.02 − 36 µm. the symbols "all" and "fa" shown in figure 9 indicate that the fitted line used the data of the compressive strength and cumulative pore volume of all mixes and mixes with fa, respectively. concrete with fa should be considered separately when investigating the correlation between drying shrinkage and cumulative pore volume because there are no liner fits with high correlation coefficients in figure 9 (a) and (c), whereas the "fa" line shows high correlation coefficients in figure 9 (b) and (d). the drying shrinkage of the concrete incorporating fa is related to the cumulative pore volume of micropores due to the pore diameter of interval divisions shown in figure 9 (a)and (c), both of which start from 0.003 µm. in other words, there is a correlation between cumulative pore volume and drying shrinkage due to the incorporation of fa in concrete. this can be explained by the fact that fa can reduce both the pore size of capillary pores and gel pores in concrete and because the shrinkage rate depends on the amount of water lost in the mesopore, in which the resistance to water loss was determined by the size of the macropore [9]. hence, the drying shrinkage of concrete with fa is relatively correlated with the pore volume of micropores and mesopores. 0.003 − b µm is the optimal interval between the drying shrinkage and the cumulative pore volume of concrete incorporating fa because the "fa" line in figure 9 (b) has the highest correlation coefficient of 334 vol. 33/2022 porosity and properties of fly ash concrete figure 9. coefficients of determination (r2) of drying shrinkage and cumulative pore volume at different interval divisions: (a) a−b µm; (b) 0 − 0.03 µm; (c) 0.02 − 0.2 µm; (d) 0.03 − 0.2 µm. 0.82. 3. conclusions both fly ash and recycled aggregates in concrete mixes were detrimental but at moderate rates; using fa in rac production was advisable, especially for percentages of fa exceeding the standard limit, since some of the particles of fa work as a pozzolanic binder ratio instead of a filler. cumulative pore volume increases with an increase of rfa replacement ratio and decreases with a decrease of the w/b ratio. concrete incorporating fa should be considered separately when investigating the correlation between cumulative pore volume and compressive strength or drying shrinkage. there is no linear correlation between compressive strength and cumulative pore volume in all interval divisions. there is a linear correlation between cumulative pore volume and drying shrinkage result from the incorporation of fa in concrete and the optimal interval division for investigating the correlation between drying shrinkage and the cumulative pore volume of concrete incorporating fa is 0.003 − b µm. the smallest pore diameter that can be measured by the mip test is 0.003 − b µm and point b is the right change point of pore diameter in a plot of the cumulative pore volume. references [1] t. c. hansen, e. boegh. elasticity and drying shrinkage of recycled concrete aggregate. journal proceedings 82(5):648-52, 1985. [2] s.-c. kou, c.-s. poon, h.-w. wan. properties of concrete prepared with low-grade recycled aggregates. construction and building materials 36:881-9, 2012. https: //doi.org/10.1016/j.conbuildmat.2012.06.060. [3] r. kurad, j. d. silvestre, j. de brito, et al. effect of incorporation of high volume of recycled concrete aggregates and fly ash on the strength and global warming potential of concrete. journal of cleaner production 166:485-502, 2017. https://doi.org/10.1016/j.jclepro.2017.07.236. [4] k. kim, m. shin, s. cha. combined effects of recycled aggregate and fly ash towards concrete sustainability. construction and building materials 335 https://doi.org/10.1016/j.conbuildmat.2012.06.060 https://doi.org/10.1016/j.jclepro.2017.07.236 h. lin, k. takasu, h. koyamada, h. suyama acta polytechnica ctu proceedings 48:499-507, 2013. https: //doi.org/10.1016/j.conbuildmat.2013.07.014. [5] m. tavakoli, p. soroushian. drying shrinkage behavior of recycled aggregate concrete. concrete international 18(11):58-61, 1996. [6] c. d. atiş, a. kiliç, u. k. sevim. strength and shrinkage properties of mortar containing a nonstandard high-calcium fly ash. cement and concrete research 34(1):99-102, 2004. https://doi.org/10.1016/s0008-8846(03)00247-3. [7] z. w. wu, h. z. lian. high performance concrete. beijing: china railway publication house, p. 5-13, 1999. [8] h.-l. ho, r. huang, w.-t. lin, et al. pore-structures and durability of concrete containing pre-coated fine recycled mixed aggregates using pozzolan and polyvinyl alcohol materials. construction and building materials 160:278-92, 2018. https: //doi.org/10.1016/j.conbuildmat.2017.11.063. [9] f. collins, j. g. sanjayan. effect of pore size distribution on drying shrinking of alkali-activated slag concrete. cement and concrete research 30(9):1401-6, 2000. https://doi.org/10.1016/s0008-8846(00)00327-6. 336 https://doi.org/10.1016/j.conbuildmat.2013.07.014 https://doi.org/10.1016/s0008-8846(03)00247-3 https://doi.org/10.1016/j.conbuildmat.2017.11.063 https://doi.org/10.1016/s0008-8846(00)00327-6 https://doi.org/10.14311/app.2022.33.0277 acta polytechnica ctu proceedings 33:277–283, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague a study on properties of concrete with dry fly ash and fly ash slurry stored with stirring xiangnan ji∗, koji takasu, hidehiro koyamada, hiroki suyama the university of kitakyushu, faculty of environmental engineering, department of architecture, japan ∗ corresponding author: b0dbb410@eng.kitakyu-u.ac.jp abstract. in the evaluation of concrete sustainability, what constitutes "sustainable" to one region may vary from another. this often leads to methodological forms of uncertainties that makes the evaluation process more complex. as such, this paper aims to quantify the effect of uncertainties in the regional context on the sustainability evaluation of concrete materials. this is carried out by quantifying the regional context through establishing a weighting scheme and then integrating the obtained weights into the sustainability analysis of concrete materials in tandem with uncertainty analysis. japan is used as a case study because although it relatively appears as a homogeneous country, its prefectures possess unique characteristics that may make the sustainability evaluation of concrete materials vary across prefectures. cluster analysis is carried out in the 47 prefectures of japan using a set of regional context indicators. five clusters are identified with varying characteristics and these are translated into different weighting schemes. the established weights are used in the sustainability evaluation of concrete materials using multi-criteria decision-making analysis. the results showed that one mix is the most sustainable for four of the clusters and a different mix is the most sustainable for the remaining cluster. when uncertainty analysis is conducted, the effect of the weights in the sustainability evaluation is explained by examining the average scores of the concrete mixes and the variance of the scores across the five clusters. this investigation facilitated the understanding of how regional differences and the uncertainties associated with it impact the evaluation of concrete sustainability. keywords: dry powder, flow, fly ash, fresh properties, slump, slurry. 1. introduction as a previous study, we invented a device to remove unburned carbon by flotation method. since the flotation method immerses the fly ash in water for treatment, it is necessary to verify the influence of the slurry of the fly ash on the concrete. therefore, in this study, fly ash from three different sources was stirred and stored for 1-28 days, and the physical changes of the three substances stored under this condition were recorded, and the freshness of the concrete using dry powder was confirmed. and check the usefulness of the slurry. 2. experimental summary 2.1. material used table 1 lists the different properties of the three fly ash. in this experiment, fly ash a, b and c from three different sources were used. they all conform to the class ii specified in jisa6201. the dry powder and the slurry are respectively used as a mixture in the concrete, and the preparation method of the slurry is: the concentration was adjusted to 60 ± 1 % using an infrared moisture meter, placed in a 90 l vessel, and stirred at a rotation speed of 60 rpm to prevent sedimentation. for the samples used for the slurry, the b and c ash were respectively stirred for 7 days, stored and reused, a ash was added on the 7th day, and samples were taken at 1, 3, 14 and 28 days. when used as a mixed material, after measuring the slurry concentration using an infrared moisture meter before putting it in, in order to meet the unit water volume of the preparation, add insufficient water, than start stir. in addition, the slurry concentration before the input showed some change after the initial adjustment, but there was no particularly significant change. table 2 shows the materials used. we use ordinary portland cement as cement, sea sand as fine aggregate and gravel as coarse aggregate. concrete mixing was carried out in a mixing chamber at a temperature of 20◦ c ± 1 using a forced biaxial mixer with a target air volume of 4.5 ± 1.0 % and a target slump of 18.0 ± 2.5 cm. in addition, regarding the external weight of 120 kg/m3 , high fluidity concrete was used, and the target slump flow amount was 50 ± 5 cm. table 3 shows the mix ratio of concrete. in this experiment, the unit water volume is 180 kg/m3, the fine aggregate rate is fixed at 45%, and the fly ash is used as an internal substitute for cement replacement, replacing 10, 20, 30%, using 120 kg/m3. ash was used as a substitute for fine aggregates and experiments were performed by external replacement. in addition, in the b and c ash, the slump and compressive strength tests were carried out at an internal replacement rate of 20%. 277 https://doi.org/10.14311/app.2022.33.0277 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en x. ji, k. takasu, h. koyamada, h. suyama acta polytechnica ctu proceedings s y m b o l t a p e s ta te b ef o re th e ex a m sl u rr y co n ce n tr a ti o n (% ) p er io d l o ss o n ig n it io n (% ) d en si ty (g / cm 3 ) b ra in sp ec ifi c su rf a ce a re a b e t sp ec ifi c su rf a ce a re a (m 2 / g ) w o rk u p te m p er a tu re (◦ c ) s lu m p fl o w (r ig h t a ft er ) s lu m p fl o w (3 0 m in ) s lu m p fl o w (6 0 m in ) s lu m p fl o w (9 0 m in ) s lu m p fl o w (1 2 0 m in ) a ir v o lu m e (i m m ed ia te ly a ft er ) (% ) a ir v o lu m e r ed u ct io n ra te (% ) a d m ix tu re (% ) p − − − − − − − − 1 8 .2 1 7 .2 1 7 .0 1 4 .4 1 3 .5 5 .0 5 0 .0 0 .4 a a − − − 1 .3 2 .3 4 0 0 0 − − − − − − − − − a 1 d -2 0 in a s 6 0 .5 1 d 2 .0 − − 3 .2 2 0 .4 1 9 .0 1 8 .8 1 7 .0 1 5 .9 4 .1 4 6 .3 0 .4 a 3 d -2 0 in s 6 0 .4 3 d 2 .2 − − 3 .1 2 0 .3 1 9 .7 1 8 .5 1 8 .0 1 8 .2 3 .5 4 0 .0 a 7 d -1 0 in d − 7 d − − − − 1 8 .4 1 8 .2 1 7 .9 1 5 .3 1 3 .3 4 .0 5 2 .5 s 5 9 .6 2 .3 − − 4 .4 1 8 .2 1 3 .7 1 4 .6 1 2 .0 8 .2 5 .0 3 8 .0 a 7 d -2 0 in d − − − − − 1 8 .7 1 7 .5 1 8 .7 1 6 .6 1 8 .3 3 .9 3 5 .9 s 6 0 .4 2 .3 − − 4 .4 1 9 .3 1 6 .8 1 7 .3 1 5 .5 1 5 .5 4 .1 3 6 .6 a 7 d -3 0 in d − − − − − 1 9 .8 1 9 .6 1 8 .5 1 7 .5 1 6 .5 3 .5 4 0 .0 0 .6 s 6 0 .2 2 .3 − − 4 .4 2 0 .3 1 9 .1 1 8 .0 1 7 .0 1 6 .0 3 .8 4 7 .4 a 7 d -1 2 0 o u t d − − − − − 5 0 .1 5 0 .0 4 4 .0 4 1 .3 3 8 .0 5 .4 4 2 .6 0 .6 (s p ) s 6 1 .2 2 .3 − − 4 .4 5 2 .6 5 1 .9 4 4 .1 3 9 .8 3 9 .6 5 .5 4 0 .0 a 1 4 d -2 0 in s 6 0 .7 1 4 d 2 .3 − − 2 .8 1 8 .8 1 6 .7 1 4 .9 9 .4 9 .0 3 .5 3 7 .1 0 .4 a 2 8 d -2 0 in s 6 1 .8 2 8 d 2 .4 − − 3 .0 2 0 .5 1 9 .5 1 8 .2 1 5 .7 1 5 .2 4 .1 5 3 .7 b b − − − 1 .6 2 .3 4 1 4 0 − − − − − − − − − b 7 d -2 0 in d − 7 d − − − − 1 7 .7 1 6 .7 1 4 .5 1 3 .3 1 2 .1 3 .5 5 4 .3 0 .4 s 6 0 .6 − − − − 1 9 .5 1 8 .2 1 7 .8 1 5 .9 1 4 .2 3 .5 4 0 .0 c c − − − 3 .0 2 .2 3 5 2 0 − − − − − − − − − c 7 d -2 0 in d − 7 d − − − − 2 0 .0 2 0 .0 1 9 .2 1 7 .3 1 6 .6 3 .6 3 0 .6 0 .4 s 6 0 .8 − − − − 1 9 .0 1 9 .0 1 6 .6 1 5 .9 1 5 .0 3 .6 3 6 .1 t ab le 1. p hy si ca l pr op er ti es of fly as h an d fr es h pr op er ti es of co nc re te . 278 vol. 33/2022 a concrete with dry fly ash and fly ash slurry figure 1. a ash s age every particle size distribution. the dry ash content of fly ash is d, the amount of slurry is s, the internal is replaced by in, and the external is replaced by out. the storage period is 1 day, 3 days, 7 days, 14 days, 28 days, respectively 1d, 3d, 7d, 14d, 28d. 2.2. experimental project the measurement items in this experiment are slump and flow. the concrete slump test was carried out in accordance with jisa1101. in order to confirm the change of the measured amount with time, the slump was recorded every 30 minutes for the first 2 hours after the start of the experiment. use fresh concrete at the beginning and end of the measurement of slump. in addition, the fresh concrete is required to be carried out in a mixing chamber at a temperature of 20◦ c ± 1◦ c. 3. experimental results and consideration 3.1. the physical properties of fly-ash as shown in table 1, for the a ash, the bet specific surface area and the strong heat loss were measured according to the age of the different preservation materials, and the former tends to increase as the material aged. previous studies have shown that the rising trend of jis ii ash has not changed and a similar trend has emerged. however, the latter was the largest at 7d and no change was observed at ages other than 7d. figure 1 shows the particle size distribution of the slurry of a ash. previous studies have reported that particles in the range of 0.1 to 1 µm tend to increase by storing the slurry, but there is no similar tendency in the jis ii seed ash of this experiment, and there is no change in particle size. figure 2. a ash slump. 3.2. concrete slump-flow table 1 shows the results of the experiment, and figure 2 shows the change in the slump of d of a and the change of s according to time. it can be seen that in addition to s10 in, the slump value at the time of placement was reduced to about 80% after 2 hours. it can be seen that within the scope of this experiment, the quality of the fresh property did not change even when used in d, s, although s10 had been reduced to about 50%. jass5 (building standard specification) stipulates that if the external temperature is lower than 25 ◦ c, it will be 120 minutes from mixing to completion, but during this period, the slump of concrete mixed with p and fly ash has the same timedependent change. as confirmed, it can be considered that the concrete containing the fly ash mixed with s can be appropriately placed. in addition, since it is impossible to clarify the time-dependent change of s10 in slump in this experiment, it will be a future research topic including re-experiment. the a, b, and c ash drop as a function of time as shown in figure 3 as a function of time. the b and c ash showed the same trend as the a ash 2 hours after the slump value. within the scope of this experiment, it was found that when the ash type equivalent to the jis ii type is used, the fresh performance of the concrete mixed with fly ash in d and s does not change. figure 4 shows the change in 20 in slump over time for s, which was stored for 1, 3 and 14 days with stirring, except for the 7 ash a ash. as a result, the slump value of s using 14d was significantly lowered. from table 1, the bet specific surface area of each s material age was confirmed, but the s value of 14d was the lowest. generally, the larger the bet specific surface area, the higher the adsorption rate of the mixture, and the fluidity becomes worse. further, in the previous study 3), fly ash having a bet spe279 x. ji, k. takasu, h. koyamada, h. suyama acta polytechnica ctu proceedings materials type physical property symbol cement ordinary portland cement density: 3.16 g/cm3 c water tap water − w fine aggregate sea sand dead density: 2.59 g/cm3 swater absorption rate: 0.76 %coarse grain rate: 2.4 performance rate: 61.2 % coarse aggregate crushed stone dead density: 2.59 g/cm3 gwater absorption rate: 1.41 %coarse grain rate: 6.9 performance rate: 56.7 % admixture high performance ae water reducing agent polycarboxylic acid ether system sp ae agent alkyl ether anionic surfactant ae table 2. materials used. figure 3. a ash, b ash, c ash, 20in slump. cific surface area of 3.6 m2/g or more was considered to be caused by deterioration of fluidity. however, in the 14d having the lowest bet specific surface area, the fluidity deteriorated due to aging. further, the value of the bet specific surface area fell within the range of the previous study, and the păof 7d was 4.4 m2/g, which showed other s ages have the same slump change. from this, it can be considered that the deterioration of the fluidity of s 14d is the most serious, and it is necessary to study this in the future. 3.3. compressive strength figure 5 shows the d and s concrete compressive strengths of a-ash. for 10in and 20in concrete, the initial strength is lower than that of p, at 91 days strength, compressive strength equivalent to p figure 4. a ash s age every 20in slump. was shown by the pozzolanic reaction of fa. 120out showed greater compressive strength than p due to fine aggregate replacement. a-ash showed a slight difference at 91 days of age, but the compressive strength tended to be higher when it was mixed with concrete in s. figure 6 shows the compressive strength of 20inch concrete with b and c ash in addition to a ash. in the case of b ash, the compressive strength was higher when mixed with d than when mixed with s, and the results were opposite to those of a and c ash. figure 7 shows the compressive strength of a ash for each age of slurry agitation storage. by compressing and storing the slurry, the compressive strength increased up to 3 days, and slightly decreased from 7 days, but the value tended to decrease as the storage period became longer. however, 280 vol. 33/2022 a concrete with dry fly ash and fly ash slurry preparation s/a unit mass (kg/m 3) % w c s g fa p 45 180 360 773 978 0 fa10in 324 768 971 36 fa20in 288 763 965 72 fa30in 252 758 958 108 fa120out 360 635 978 120 table 3. mix proportion. acceleration symbol condition w/b (%) neutralization rate coefficient (mm/ √ week) p − 50 3.70 a7d-10in d 4.76s 4.24 a7d-20in d 4.72s 4.69 a7d-30in d 7.13s 6.96 a7d-120out d 38 4.98s 4.45 table 4. neutralization rate coefficient. figure 5. a ash compressive strength. it showed almost the same compressive strength as p at all ages. as in the case of the slump test, it was clarified that, even within the range of this experiment, the performance was equivalent to that of d even when stored under stirring with s. 3.4. accelerated neutralization figure 8 shows the results of the accelerated neutralization test for concrete mixed with d and s in aăash, and table 4 shows the neutralization rate coefficients figure 6. a, b, c ash 20in compressive strength. for each. compared with p, the carbonation of concrete containing fa was more likely to increase in both d and s. it is considered that the alkali component in the concrete decreased due to the decrease in the amount of cement, and the progress of neutralization was accelerated. further, in the accelerated neutralization test, a long-term increase in strength due to a volcanic ash reaction was observed in high concentrations of carbon dioxide from an early age, this is probably because the densification of the or281 x. ji, k. takasu, h. koyamada, h. suyama acta polytechnica ctu proceedings figure 7. 20-in compressive strength for each age of a-ash s. figure 8. a ash accelerated neutralization. ganization was not considered. 3.4.1. drying shrinkage figure 9 shows the results of the concrete drying shrinkage test mixed with d and s of the ash. both d and s showed the same tendency in all formulations from the initial drying shrinkage value to about 4 months. although slightly smaller than p, the dry shrinkage value tended to be smaller when fa was mixed in all the formulations, indicating the effect of reducing the dry shrinkage of fa. 3.5. pore structure figure 10 shows the change in pore size distribution of a ash 20in, 120out, b ash and c ash 20in with time, divided into four ranges: 0.003µm 0.05µm, 0.05µm 0.5µm, 0.5µm 5µm, 5µm 100µm. it can figure 9. a ash drying shrinkage. be seen that the total pore volume tends to decrease with age for most formulations and ash types. in addition, it can be seen that in all mixtures and ash types, the value of the number of pores in the range of 0.5 to 5 µm and 5 to 100 µm hardly changes as the material ages. in the range of pore diameters from 0.003 to 0.05µm, a-ash 20in showed a slight increase tendency, but b-ash and c-ash did not show any increase or decrease. it can be considered that due to the hydration reaction, large pores with a diameter ranging from 0.05 µm to 0.5 µm transition to small diameters, and the number of pores has not changed significantly, and a similar trend has been observed in previous studies. in addition, it can be seen that the pore volume decreases most in the range of 0.05ăµm to 0.5 µm with the age of the wood, depending on the composition and ash type. from the above, it is inferred that the pore volume in the range of 0.05 µm to 0.5 µm is related to the compressive strength. figure 11 shows the relationship between the amount of pores in the range of 0.05 µm to 0.5 µm and the compressive strength. it can be seen that the relationship between the range of 0.05 µm to 0.5 µm and the compressive strength is high in all cases, regardless of the state, type, and composition of the fa. there is a possibility that the compressive strength is increased by decreasing the amount of pores in this range due to the hydration reaction and making the structure denser. previous studies have shown aăcorrelation with compressive strength in the range of 0.03 µm to 0.3 µm, confirming a similar relationship with pore volume. also, when comparing the amount of pores with the same level of compressive strength, if the amount of pores is the same, the compressive strengths are almost the same. therefore, it was shown that the pore volume could be inferred from the compressive strength even if the state of fa was different. 282 vol. 33/2022 a concrete with dry fly ash and fly ash slurry figure 10. pore size distribution. figure 11. compressive strength and pore volume (0.05 0.5µm). 4. summary concrete slumps showed no difference within 2 hours regardless of the fa condition. it is considered that there is no difference in slump properties within the range of this experiment even when s is continuously stirred and stored. regarding the compressive strength, regardless of the state of fa, the long-term strength of 10in and 20in showed the same compressive strength as that of p, and it was shown that the compressive strength manifestation may be higher when s is mixed. regarding the neutralization rate coefficient, the result was larger when fa was mixed regardless of the fa state. the neutral acceleration coefficient tended to be smaller when s was mixed. from the relationship between the compressive strength and the amount of pores, it was shown that the compressive strength could be increased by decreasing the amount of pores in the range of 0.05 to 0.5 µm regardless of the state of fa. the drying shrinkage shows almost the same tendency regardless of the fly ash condition. in addition, it was confirmed that the drying shrinkage reducing effect of fly ash showed the same tendency regardless of the state. 283 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2021.31.0041 acta polytechnica ctu proceedings 31:41–44, 2021 © 2021 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the assessment of priorities in the construction of new high-speed railway lines in czech republic by using the stem method pavel purkart∗, jan kruntorád, lukáš týfa czech technical university in prague, faculty of transportation sciences, department of transportation systems, horská 3, 128 03 prague 2, czech republic ∗ corresponding author: purkapav@fd.cvut.cz abstract. the aim of the paper is to set priority of the fast connection sections ("rychlá spojení" in czech, abbreviation "rs") building mentioned above by using the stem method. the method is used to select projects with the maximum benefit if the budget is limited. there has not been set a limit for individual rs sections as a limitation of financial costs. the decisive factor is the order in which individual rs sections should be built on the basis of evaluation criteria. keywords: high-speed railway lines (hsl), the fast connections (rs), czech republic, mathematic model, the stem method, project choice, evaluation criteria. 1. introduction and goals today a great effort is being made in the czech republic to prepare the hsl network in relation to the previously developed rs concept. now, the czech railway network does not have any line in hsl parameters, which is not acceptable in the long term – with regard to sustainable development of the territory, taking into account the situation not only in the neighbouring western european states, but also global development [1]. the main directions of the fast railway network have been defined [2]: • rs1 praha (prague) – brno – ostrava – [poland], • rs2 brno – [slovakia/austria], • rs3 praha – plzeň (pilsen) – [germany] (western direction – münchen), • rs4 praha – [germany] (northern direction – dresden) with possible branch (most – chomutov, žatec, louny), • rs5 praha – hradec králové/liberec – [poland]. the implementation of the rs1, rs2 and rs4 sections is monitored as a priority. in these cases, the správa železnic (the main infrastructure manager of czech railway network) is already carrying out project preparation. the aim of this article is thus to open up the controversy as to whether the above-mentioned sections in terms of benefits should be actually built first, respectively whether there is a more appropriate order of implementation. the current design state is schematically shown in figure 1. 2. understanding and solving the issue the main aim of the paper is to set priorities in the construction of high-speed railway lines in the czech republic using the stem method (step method). the use of the stem method is defined in the following chapter. the method is used for project selection due to limited funding opportunities while maximizing the benefits of their implementation. in this case, the method was slightly modified (the method mandatory used for economic problems was modified for a purpose for maximization benefits within economic costs at first). 3. solution of the issue 3.1. method stem the stem method can solve linear mathematical problems with more purpose functions. the aim of this method is to find compromise solutions, whose realizations should bring most benefits. the main principle of the method is the calculation of purpose function ideal values for individual cases. this calculation is followed by minimizing compromise solution deviation from ideal purpose function values. the basic of the method is an interactive procedure of searching the compromise solution. benefit of the stem method is, that there is only minimal need of communication between a submitter and a solver (comparing to another methods). the scale method for individual criterions is set by calculation. the submitter must decide whether the result of the calculation is acceptable for him or not. so, the method consists of calculation and decision-making process. the calculation is stopped, if submitter finds the result acceptable, otherwise the solver must be informed by submitter in order to change the criteri41 https://doi.org/10.14311/app.2021.31.0041 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en p. purkart, j. kruntorád, l. týfa acta polytechnica ctu proceedings figure 1. suggestion of rs, source: správa železnic [3]. ons or their numbers, the whole calculation is made again [3]. the stem method consists of following steps: (1.) solver calculates optimal solution for individual criterions (purpose function) separately. the number of calculations fits the number of criterions. (2.) solver calculates the scales of individual criterions (χ;xj ) according to the formula (1): wi = zii − min i=j...k zij zii α√ n∑ i=1 c2ij (1) where: zij – element of optimalization criterions values matrix for optimalization in individual optimalization criterion (zij is the value of optimalization criterion j = 1, . . . ,k in case of optimalization according to the criterion i = 1, . . . ,k), cij – element of the price matrix – element of individual optimalization criterion coefficients matrix. value α comes from the equation (2): k∑ i=1 zii − min i=j...k zij zii α√ n∑ i=1 c2ij = 1 (2) in reality we have to calculate coefficient α value first and then count the scales of individual criterions. if the scale fits the constraint wi > 0 for more criterions, the solver adds new variable and solves the model with new optimalization criterion (3). min f(x,d) = d (3) there is a form (4) for variable d: d = max i=j...k { wi ( zii − ∑ j∈j cijxj )} (4) we have to implement constraint (5) for correct calculation: wii ( zii − ∑ j∈j cijxj ) ≤ d (5) if constraint wi > 0 fits for only one value i = 1, . . . ,k, solver cen simplify the constraint (5) to (6): min f(x) = k∑ i=1 wii ( zii − ∑ j∈j cijxj ) (6) (3.) solver presents the results to the submitter. the submitter can modify the criterions or add/remove some of them, if he does not find the results acceptable – but we haven’t to do that, if the first solution should be the best. solver goes back to step 2. (4.) solver has found compromise solution, if the submitter of satisfied with the result. the solution is optimal, if the value d = 0 is reached. 42 vol. 31/2021 the assessment of priorities in the construction 3.2. calculation according to the stem method and result for the purposes of the calculation, a total of six logical rs sections were identified: (1.) praha – brno, (2.) brno – ostrava – [poland], (3.) brno – [slovakia/austria], (4.) praha – plzeň – [germany], (5.) praha – ústí nad labem – [germany] (northern direction – dresden) with branch, (6.) praha – hradec králové/liberec – [poland]. the following evaluation criteria were used to determine priorities in the construction of rs sections: • average coefficient of reduced travel times in chosen connections: the research team (authors of the articles + colleagues from the department) determined rs sections, including international ones, for which the rs construction is crucial. for these sections there was calculated the number of shorter travel times compared to the current situation and in the case rs section is realized. then these numbers were distributed on individual rs sections and the average for these sections were calculated, • number of redirected long-distance lines expressed as an importance: three categories were set for each evaluated section based on the evaluation below: 1: up to 10 redirected long-distance routes in both directions per day, 2: up to 30 redirected long-distance routes in both directions per day, 3: over 30 redirected long-distance routes in both directions per day. in this criterion there was monitored the possible placement on the rs network of the current longdistance lines, especially close to agglomerations (prague, brno, ostrava), where the selected sections of the railway network are already overloaded, • routing difficulty: the average was calculated based on the questioning an expert team of four experts. the issue of routing was considered broadly (not only regarding natural barriers, but also routing through urban areas, etc.) and the logic of values is divided into these categories: 1: very difficult routing, 1,5:difficult routing, 2: medium difficult routing, 2,5:routing without major problems, 3: simple routing. • road traffic redirection potential: for all rs sections, a section on the road network was determined, from which the decisive load value in vehicles/24 hours was taken. the source of data is national traffic census 2016. in all cases, the section on a parallel highway network with logical minimum value of traffic volumes was taken into account, so this criterion represents the minimal potential for traffic redirection from roads. • agglomeration potential: the research team set agglomerations, including foreign ones, on which the built rs sections will have a direct influence. for all agglomerations the number of inhabitants was found and then their relation to individual rs sections was evaluated. for example, for pilsen agglomeration the rs3 section is crucial, but the rs2 section has a minimal importance – this logic was used to evaluate all agglomerations regarding a number of inhabitants. the importance of an agglomeration for a given section was multiplied with its number of inhabitants and the final values were cumulated into individual sections. all criteria have been compiled in such a way that as their value increases, their benefits are maximized. the specific values of evaluation criteria are summarized in table 1. the stem method is a very interactive method because it calculates the weight criteria for every mathematical model. there were two mathematical models considered in the calculation (the first one to select two projects out of six and the second one to select four projects out of six), therefore weights were determined twice. these weights are shown in table 2. in both models, the highest criteria weights were assigned to number of redirected long-distance lines expressed as an importance and routing difficulty, both above 30%. the average coefficient of reduced travel times in chosen connections has a value oscillating around a weight of 20%. on the other hand, in both calculations there were assigned very low values for road travel redirection potential and agglomeration potential, these weights are lower than 5%. the calculation was performed for both versions of the model, and in neither case d = 0 was achieved. it shows that compromise solutions, not optimal ones, have been achieved. the two top prioritized high-speed railway line sections have been set by the stem method: • praha – brno, • brno – [slovakia/austria]. these two sections have middle priority: • brno – ostrava – [poland], • praha – plzeň – [germany]. the rest of sections was described as a low building priority: • praha – ústí nad labem – [germany with branch], • praha – hradec králové/liberec – [poland]. 43 p. purkart, j. kruntorád, l. týfa acta polytechnica ctu proceedings section of railway line rs average coefficient of reduced travel times in chosen connections [-] number of redirected longdistance lines expressed as an importance [-] routing difficulty [-] road traffic redirection potential [vehicles per day in thousands in chosen point of a road network] agglomeration potential [-] praha – brno 2, 13 3 1, 6 39 22 brno – ostrava – pl 1, 85 2 2, 5 22 15 brno – sk/a 1, 85 2 3, 0 23 21 praha – plzeň – d 1, 73 3 1, 6 31 13 praha – ústí n. l. – d 2, 23 2 1, 5 24 20 praha – hk/lbc – pl 2, 39 1 1, 9 23 12 table 1. specific values of evaluation criteria, source: authors. criterion weight criterion for selecting 2 projects weight criterion for selecting 2 projects average coefficient of reduced travel times in chosen connections [-] 0.181 0.251 number of redirected longdistances lines expressed as an importance [-] 0.407 0.372 routing difficulty [-] 0.361 0.337 road traffic redirection potential [vehicles per day in thousands in chosen point of a road network] 0.024 0.013 agglomeration potential [-] 0.027 0.028 table 2. weight criteria set by stem method for cases of selecting 2 and 4 projects, source: authors. 4. conclusion it is obvious from the calculation, according to evaluation criteria, that the top priorities of construction have sections praha – brno and brno – [slovakia/austria]. while the plans of správa železnic (the main infrastructure manager of czech railway network) counts with the section praha – ústí nad labem – [germany], result of this paper prefers a construction of the section praha – plzeň – [germany]. it is also possible to assume that in the section prague ústí nad labem – [germany], the railway already uses its potential to a considerable extent, so its construction was not identified as a priority by the model. it is evident that the significant influence on the calculation according to the stem method had the number of long-distance transport lines transferred, which in the case of the route through plzeň is relatively significant (lines ex6, r16 and r26). this factor apparently helped significantly to increase the priority of construction of the rs through plzeň than through ústí nad labem. acknowledgements all the facts presented in the article are based on the results of research on ctu faculty of transportation sciences, department of transport systems. this work was supported by the grant agency of the czech technical university in prague, grant no. sgs18/150/ohk2/2t/16 railway track parameters for transportation service optimization and no. sgs20/138/ohk2/2t/16 design and optimal use of railway infrastructure parameters. references [1] p. purkart. optimal cooperation of public transport segments. praha: defense date 2019-12-06. doctoral minimum. ctu fts. department of transportation systems. supervised by týfa, l. [2] high-speed rails in the czech republic. [online], praha: správa železnic, 2020 [cit. 2020-02-22]. available from: https://www.szdc.cz/vrt/. [3] d. teichmann, m. dorda. comparison of two selected methods in evaluating of investments in transport infrastructure. in: finance and performance of firms in science, education and practice : proceedings of the 7th international scientific conference : april 23-24, 2015, zlín, czech republic. zlín: tomas bata university in zlín, 2015. pp. 1524-1536. isbn 978-80-7454-482-8. 44 https://www.szdc.cz/vrt/ acta polytechnica ctu proceedings 31:41–44, 2021 1 introduction and goals 2 understanding and solving the issue 3 solution of the issue 3.1 method stem 3.2 calculation according to the stem method and result 4 conclusion acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0141 acta polytechnica 27(0):141–144, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app indentation size effect in high pressure torsion processed high entropy alloy petr haušilda, ∗, jakub čížekb, jaroslav čecha, jiří zýkac, hyoung seop kimd a czech technical university in prague, faculty of nuclear sciences and physical engineering, department of materials, trojanova 13, 120 00 praha, czech republic b charles university in prague, faculty of mathematics and physics, department of low temperature physics, v holešovičkách 2, praha 8, cz-18000, czech republic c ujp praha a.s., nad kamínkou 1345, 156 00, praha, czech republic d postech, department of materials science and engineering, pohang, 790-784, south korea ∗ corresponding author: petr.hausild@fjfi.cvut.cz abstract. high entropy alloy hfnbtatizr in as cast conditions and after high pressure torsion straining was characterized by nanoindentation. the length-scale dependent material response (indentation size effect) was characterized by indentation at various indentation depths. hardness dependence on the characteristic length (depth of penetration) indicated decomposition of disordered high entropy alloy in the as cast sample, which probably occurred during slow cooling after casting. subsequent severe plastic deformation by high pressure torsion led on the other hand to the short-range disorder of (originally partially decomposed as cast) structure. further hardening was generated during high pressure torsion by the mechanisms of grain refinement and increasing dislocation density. keywords: high entropy alloys, high pressure torsion, indentation size effect, nanoindentation. 1. introduction in high entropy alloys (heas), the configurational entropy of a multicomponent solid solution phase is maximized so that the gibbs energy of random solution may be lower than that of possible intermetallic phases. generally, to achieve high entropy of mixing, the alloys contain typically five or more major elements in equimolar concentrations [1]. the heas show promising mechanical properties in applications ranging from (high temperature) structural to biocompatible materials [2, 3]. mechanical properties of such alloys can further be improved by grain refinement especially by severe plastic deformation. however, studies of ultrafine grained heas are rather scarce in the literature. an increase of strength with decreasing grain size was achieved through processing by high pressure torsion (hpt) in the probably most investigated hea cantor alloy i.e., equiatomic cocrfemnni with face-centered cubic (fcc) structure [4]. fewer attempts were made to process in such a way heas with body-centered cubic (bcc) structure. recently hfnbtatizr bcc hea was successfully nanostructured by hpt straining [5, 6]. it was reported that grain refinement by hpt resulted in a significant enhancement of the strength of this bcc hea, keeping excellent ductility during room temperature straining. nevertheless, there is still a lack of information about the development of microstructure and physical properties of this refractory metal hea subjected to severe plastic deformation processing. recent investigations [5, 6] revealed that thermodynamically stable system of hfnbtatizr alloy at room temperature is a mechanical mixture of zr, hf rich hcp phase and ta, nb rich bcc phase. recrystallization annealing after cold-rolling at 800 °c led to the transformation of the single-phase bcc material into a two-phase structure composed of bcc1 and bcc2 phases with lattice parameters a1 ∼ 0.3427 nm and a2 ∼ 0.3338 nm, respectively [7]. chemical analysis revealed that the matrix was slightly enriched in hf and zr, while the precipitated second-phase particles were on the other hand slightly enriched in ta and nb. the decomposition of the solid solution (and formation of short range order) after long-term annealing obviously leads to the deterioration of mechanical properties (loss of ductility and decrease of strength). the difference in hardness of both phases is relatively small and both are softer than the random (disordered) solid solution [6]. on the other hand, considerable contribution to the solid solution strengthening can arise from atomic size misfit (phase separation on the nano-meter scale) which is provoked by the high density of vacancies introduced by hpt [5]. the aim of this work was clarification of the relationship between phase (de)composition, microstructure, lattice defects, and length-scale-dependent material response of hfnbtatizr hea in as cast conditions, after solution (disordering) thermal treatment and hpt straining. the length-scale-dependent material response was characterized by indentation at various 141 http://dx.doi.org/10.14311/app.2020.27.0141 http://ojs.cvut.cz/ojs/index.php/app p. haušild, j. čížek, j. čech et al. acta polytechnica as cast annealed 1/4 turn 1 turn 15 turns εpl [1] 0 0 5 20 300 h0 [mpa] 4049 4438 4947 5717 5953 h∗ [nm] 45.7 76.3 49 33.5 29.3 table 1. list of the samples equivalent plastic strain εpl, identified characteristic lengths h∗ and hardnesses in the limit of infinite depth h0. indentation depths. the contributions of different mechanisms were attributed to distance between dislocation pinning defects (length scale) and are reported here. 2. experimental details 2.1. material the alloy of equimolar composition hfnbtatizr was prepared by vacuum arc melting from pure hf, nb, ta, ti and zr metals and cast into water cooled cu crucible. in order to ensure proper mixing, the ingot was flipped and remelted eight times. disc of 20 mm diameter and 1 mm thickness were hpt strained at room temperature and pressure of 2.5 gpa. straining was performed to 1/4 turn, 1 turn and 15 turns of the piston (see fig. 1). corresponding equivalent plastic strains εpl (at the characterized areas) are 5, 20 and 300 (see table 1). more details can be found in ref. [5]. one non-deformed (as cast) sample was also annealed for 2 mins at 1600 °c (followed by rapid cooling) for reasons described in section results and discussion. sample figure 1. scheme of high pressure torsion straining. 2.2. nanoindentation surface of samples was polished by standard metallographical procedure with final step made using 0.04 µm colloidal silica to avoid the surface layer affected by mechanical grinding and polishing. nanoindentation measurements were performed using an anton paar nht2 nanoindentation tester with berkovich diamond indenter using instrumented indentation technique [8, 9]. in order to obtain the mechanical response from different depth, so called continuous multi cycle (cmc) indentations with increasing load were performed varying the maximum load at each cycle from 1 mn up to 500 mn. this type of load cycle allows characterization at different load (depth) levels at the same position (see e.g. refs. [10, 11]). loading time was 10 s per each cycle for cmc, followed by 5 s hold at the maximum load and unloading time 10 s per each cycle. the results (indentation hardness) were evaluated according to the iso 14577 standard. the indentations were made on sample surface in the position corresponding to the half radius, i.e. in the location of most homogeneously distributed deformation after hpt straining. at least 20 cmc indentations were performed for each material state (deformation condition). 2000 3000 4000 5000 6000 7000 8000 9000 0 1000 2000 3000 h (m p a ) h (nm) 15 1 1/4 as cast figure 2. indentation size effect in as cast sample and samples hpt strained to 1/4 turn, 1 turn and 15 turns. 3. results and discussion results of hardness measurements as a function of penetration depth for as cast sample and samples hpt strained from 1/4 turn to 15 turns are shown in fig. 2. it can be seen in this figure that hardness increases with increasing strain and that all samples show strong indentation size effect in sub-micron penetration depths. using nix nad gao approach [12] based on the geometrically necessary dislocations, the 142 vol. 27/2020 indentation size effect in hpt processed high entropy alloy depth dependence of the hardness h can be related to the characteristic length h∗ through the relation: h h0 = √ 1 + h∗ h (1) where h is a given depth of penetration, and h0 is the hardness in the limit of infinite depth (∼ macroscopic hardness). depth dependence of the hardness is plotted according to eq. 1 for different levels of straining in fig. 3 and the parameters h0 and h∗ obtained by least square fit are listed in table 1. 0 10 20 30 40 50 60 0 0.001 0.002 0.003 0.004 0.005 0.006 0.007 h 2 ( g p a 2 ) m il li o n s 1/h (nm-1) as cast 15 turns 1 turn 1/4 turn figure 3. depth dependence of the hardness plotted for as cast sample and samples hpt strained from 1/4 turn to 15 turns according to eq. 1. as can be expected, the hardness in the limit of infinite depth h0 is increasing with the increasing level of hpt straining (hpt leads to the strain hardening [13]). on the other hand, the characteristic length h∗ is decreasing with the increasing level of deformation. it is to note that the values of characteristic lengths h∗ are quite low, which indicates fine microstructural features involved in the dislocation pinning process. according to nix and gao [12], the characteristic length h∗ depends on the hardness in the limit of infinite depth h0 as follows: h∗ ∼ 1 h02 (2) fig. 4 shows that characteristic depth h∗ dependence on the hardness in the limit of infinite depth h0 fulfils relation (2) for hpt deformed samples, but the as-cast sample clearly does not follow this relation. as mentioned in the introduction, the anomalous behavior of the as cast sample is likely caused by the decomposition of disordered (high entropy) bcc solid solution into two bcc1 and bcc2 phases. such short-range atomic redistribution is hard to be characterized by conventional techniques such x-ray diffraction or transmission electron microscopy as the variation of lattice parameter is very small 0.00 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0 20 40 60 80 100 1 /h 0 2 ( g p a -2 ) h* (nm) as cast annealed1/4 turn 1 turn 15 turns figure 4. characteristic depth dependence of the hardness in the limit of infinite depth for as cast sample, annealed sample and samples hpt strained from 1/4 turn to 15 turns. 2000 3000 4000 5000 6000 7000 8000 0 500 1000 1500 2000 2500 h (m p a ) hc (nm) as cast annealed figure 5. comparison of depth dependence of the hardness in the as cast and annealed samples. and the size of domains is in the order of tens of nm (see parameter h∗). therefore, the as cast sample was annealed at (high) temperature of 1600 °c with subsequent rapid cooling in order to preserve the random solid solution and subsequently characterized by nanoindentation. the results of hardness measurements in as cast and annealed state are compared in fig. 5. as can be seen in table 1, the parameter h∗ significantly increased after annealing, which is consistent with the hypothesis of short-range order disappearance. the tendency given by the relation (eq. 2) is now in good agreement for both annealed and hpt strained samples (fig. 4). thus, similarly to high temperature annealing the hpt led to the short-range disorder of (originally partially decomposed) as-cast sample (severe plastic deformation leads to an extension of solid solubility [14]). subsequent hardening in 143 p. haušild, j. čížek, j. čech et al. acta polytechnica hpt strained samples is related rather to grain (cell) refinement and increasing dislocation density as observed in ref. [5] using transmission electron microscopy. 4. conclusions the results of the present study can be summarized as follows: • high pressure torsion led to a significant (and progressive) hardening of high entropy hfnbtatizr alloy. • both as cast and hpt strained samples showed length-scale dependent material response (indentation size effect). • as cast sample presented different dependence of characteristic length h∗ on the hardness in the limit of infinite depth h0 than hpt deformed and annealed samples, which indicates short-range decomposition of disordered high entropy alloy. • it can be concluded that severe plastic deformation by hpt produced the short-range disorder of originally partially decomposed as cast structure, which was followed by hardening due to the grain refinement and increasing dislocation density intorduced by hpt process. • annealing of as cast sample at 1600 °c with subsequent rapid cooling in order to stabilize the disordered structure led to the same dependence of characteristic length on the hardness in the limit of infinite depth as in the case of hpt strained samples. acknowledgements this research was carried out in the frame of the projects cz.02.1.01/0.0/0.0/16_019/0000778 (centre of advanced applied sciences) and 17-17016s (czech science foundation). references [1] m.-h. tsai, j.-w. yeh. high-entropy alloys: a critical review. materials research letters 2(3):107–123, 2014. doi:10.1080/21663831.2014.912690. [2] d. miracle, o. senkov. a critical review of high entropy alloys and related concepts. acta mater 122:448 – 511, 2017. doi:10.1016/j.actamat.2016.08.081. [3] j. chen, x. zhou, w. wang, et al. a review on fundamental of high entropy alloys with promising high-temperature properties. j alloy comp 760:15 – 30, 2018. doi:10.1016/j.jallcom.2018.05.067. [4] a. heczel, m. kawasaki, j. lábár, et al. defect structure and hardness in nanocrystalline cocrfemnni high-entropy alloy processed by high-pressure torsion. j alloy comp 711:143–154, 2017. doi:10.1016/j.jallcom.2017.03.352. [5] j. čížek, p. haušild, m. cieslar, et al. strength enhancement of high entropy alloy hfnbtatizr by severe plastic deformation. j alloy comp 768:924–937, 2018. doi:10.1016/j.jallcom.2018.07.319. [6] b. schuh, b. völker, j. todt, et al. thermodynamic instability of a nanocrystalline, single-phase tizrnbhfta alloy and its impact on the mechanical properties. acta mater 142:201–212, 2018. doi:10.1016/j.actamat.2017.09.035. [7] o. n. senkov, s. l. semiatin. microstructure and properties of a refractory high-entropy alloy after cold working. j alloy comp 649:1110–1123, 2015. doi:10.1016/j.jallcom.2015.07.209. [8] w. oliver, g. pharr. an improved technique for determining hardness and elastic-modulus using load and displacement sensing indentation experiments. j mater res 7(6):1564–1583, 1992. doi:10.1557/jmr.1992.1564. [9] w. oliver, g. pharr. measurement of hardness and elastic modulus by instrumented indentation: advances in understanding and refinements to methodology. j mater res 19(1):3–20, 2004. doi:10.1557/jmr.2004.0002. [10] p. haušild, a. materna, l. kocmanová, j. matějíček. determination of the individual phase properties from the measured grid indentation data. j mater res 31(22):3538–3548, 2016. doi:10.1557/jmr.2016.375. [11] p. haušild, j. čech, a. materna, j. matějíček. statistical treatment of grid indentation considering the effect of the interface and the microstructural length scale. mech mater 129:99 – 103, 2019. doi:10.1016/j.mechmat.2018.11.006. [12] w. nix, h. gao. indentation size effects in crystalline materials: a law for strain gradient plasticity. j mech phys solids 46(3):411–425, 1998. doi:10.1016/s0022-5096(97)00086-0. [13] a. p. zhilyaev, t. g. langdon. using high-pressure torsion for metal processing: fundamentals and applications. progress in mater sci 53(6):893 – 979, 2008. doi:10.1016/j.pmatsci.2008.03.002. [14] o. senkov, f. froes, v. stolyarov, et al. microstructure of aluminum-iron alloys subjected to severe plastic deformation. scripta mater 38(10):1511 – 1516, 1998. doi:.10.1016/s1359-6462(98)00073-6. 144 https://doi.org/10.1080/21663831.2014.912690 https://doi.org/10.1016/j.actamat.2016.08.081 https://doi.org/10.1016/j.jallcom.2018.05.067 https://doi.org/10.1016/j.jallcom.2017.03.352 https://doi.org/10.1016/j.jallcom.2018.07.319 https://doi.org/10.1016/j.actamat.2017.09.035 https://doi.org/{10.1016/j.jallcom.2015.07.209} https://doi.org/{10.1557/jmr.1992.1564} https://doi.org/{10.1557/jmr.2004.0002} https://doi.org/{10.1557/jmr.2016.375} https://doi.org/10.1016/j.mechmat.2018.11.006 https://doi.org/{10.1016/s0022-5096(97)00086-0} https://doi.org/10.1016/j.pmatsci.2008.03.002 https://doi.org/.10.1016/s1359-6462(98)00073-6 acta polytechnica 27(0):141–144, 2020 1 introduction 2 experimental details 2.1 material 2.2 nanoindentation 3 results and discussion 4 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0377 acta polytechnica ctu proceedings 33:377–382, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague sustainable concrete production with recycled concrete wash water beneficiated with co2 sean monkmana, ∗, vincent meyerb a carboncure technologies, 42 payzant av, dartmouth ns, canada b holcim innovation center, 95 rue du montmurier, 38070 saint-quentin-fallavier, france ∗ corresponding author: smonkman@carboncure.com abstract. a significant amount of wastewater is generated through the cleaning of equipment utilized within the ready mixed concrete production cycle. the reuse of concrete wash water as mix water is limited by the negative material performance impacts associated with the suspended solids; the effects are exacerbated with increasing solids contents and water aging. a novel carbon dioxide treatment to allow the use of high solids wash water (specific gravity 1.10) as mix water was examined. seven batches of concrete were produced and compared: a reference mix, two batches with untreated wash water and four batches with co2 treated wash water. the carbon dioxide treatment mineralized co2 at 28% by mass of the treated solids. acceptable concrete was produced through adjusting admixtures for workability. the compressive strength at 1, 7, 28 and 56 days was improved relative to both the reference and the concrete produced with untreated wash water. the suspended solids containing mineralized co2 served as a viable cement replacement. the avoided cement and bound carbon dioxide served to lower the carbon impact of the concrete by about 14%. the approach allows three waste streams (co2, wash water and wash water solids) to be reused to produce more sustainable concrete. keywords: co2 utilization, concrete, sustainability. 1. introduction awareness of the environmental impacts of the cement and concrete industries, combined with increasing global infrastructure demands, has resulted in a greater emphasis on improving concrete sustainability. strategic roadmaps and technology assessments have recognized the shortcomings of traditional levers for reducing the carbon impact of the industries and the need for emerging technologies [1]. action is require throughout the supply chain; cement producers, concrete producers, structural engineers and construction companies each have roles to play in the pursuit of a net zero carbon vision [2]. within the context of emerging technologies lies the concept of co2 utilization [3]. in this approach, carbon dioxide is used directly or as a feedstock to produce valuable carbon-containing products. one use is for the manufacture of construction materials; in contrast to other utilization pathways, such as the production of fuels, the carbon dioxide is mineralized and permanently removed from the atmosphere [4, 5]. construction has been identified as the largest market opportunity for the co2 utilization field with technologies in concrete and aggregate production offering a potential to utilize 5.0 gt co2 and generate $550b per year by 2030 [6]. the present work concerns an investigation into the use of carbon dioxide to beneficially treat concrete wash water for its reuse as mix water. the production of ready mixed concrete involves the washing of equipment whether it be a central mixer or individual trucks. after the concrete has been discharged the mixer may be washed out to prepare it for a new load of concrete. the water that is washed from the truck is laden with very fine sand and cementitious materials. the handling and disposal of this wastewater represents a significant operational, regulatory and logistical burden for the concrete producer. the most common approaches to address wash water are clarification and discharge, reclaimers, hydration stabilizing admixtures or no action at all. ideally the wash water can be recycled into new concrete but performance issues often arise when it is used as mix water in lieu of fresh water [7]. the approach can lead to concrete with increased water demand (and lower compressive strengths) and undesirable set acceleration. both of these impacts worsen with the age of water meaning that the performance outcomes are sensitive to the "freshness" of the wash water. in practice, the solution has been to specify that mix water must have less than 50,000 ppm solids [8]. in cases where wash water is used in new concrete it is often merely diluted with fresh water to reduce the negative impacts. while the practice of clarifying water and disposing solids is the current status quo, the industry is being directed towards zero discharge practices and ready mixed producers are looking to decrease their water usage and carbon footprints [9]. the present study concerns the co2 treatment of 377 https://doi.org/10.14311/app.2022.33.0377 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en sean monkman, vincent meyer acta polytechnica ctu proceedings batch control nt nta ct cta ctp ctap ww treatment n/a none none co2 co2 co2 co2 cement (kg/m3) 267 267 267 267 267 240 240 limestone(kg/m3) 63 33 33 30 30 57 57 ww solids (kg/m3) 0 30 30 33 33 33 33 total powder (kg/m3) 330 330 330 330 330 330 330 pce (%/b) 0.34 0.34 0.34 0.34 0.34 1.15 0.77 dispersant (%/b) 0 0 0.66 0 0.74 0 0.74 relative binder 100 91 % 91 % 90 % 90 % 90 % 90 % table 1. mix design of concrete produced during test program. ww refers to wash water. ww solids are the suspended solids present in the wash water that are incorporated into the concrete mix. pce refers to a polycarboxylate plasticizing admixture. admixture dosages are in % with respect to the binder content. high solids concrete wash water for reuse, without dilution, as concrete mix water. carbon dioxide is readily mineralized into calcium carbonate upon reaction with tricalcium or dicalcium silicate [10, 11] or calcium hydroxide [12], wherein the carbon dioxide is permanently removed from the atmosphere it was studied what impact the carbon dioxide would have on the use of wash water as mix water and whether performance-enhanced concrete could be produced. 2. experimental a simulated concrete wash water was produced by mixing 7.7 kg of cement with 18 litres of water. the slurry was mixed in a drum mixer at 24 rpm for 3 hours to simulate a production/delivery/washing cycle. the slurry was then added to a treatment vessel. the total water in the reactor vessel was brought up to 50 litres by adding 7 litres of water through wash out of the drum mixer and a final 25 litres of water. the slurry was designed to have a specific gravity of 1.10 or more than three times the nominal optional limit for concrete mix water that appears in many specifications such as astm [8]. the supply of slurry was maintained under agitation in a barrel adapted as a reactor vessel. an immersible pump served to circulate the slurry through a conduit that was routed up and out of the slurry before bending back down to the vessel bottom whereupon the pumped slurry was ejected. wash water subjected to a co2 treatment received an injection of carbon dioxide gas at 600 litres per hour. the injection of co2 gas was integrated into the base of the conduit where it was connected to the exit of the pump. the conduit, approximately 1.2 meters in total length, served to increase the potential residence time of gas inside the slurry given that any bubbles would rise through the slurry via buoyancy. three different batches of slurry were made (one untreated, two treated) and each batch of slurry was used to make two batches of concrete. both treated and untreated wash water slurries were kept agitated in a minimally rotating drum mixer through to 24 hours prior to use in concrete. samples of the treated slurry were taken periodically throughout the treatment cycle which concluded when the ph had declined from an originally elevated level to a neutral level of about 7.0. the carbon uptake and bound water of the cement slurry solids was determined through thermogravimetric analysis. a second assessment of the bound carbon was completed using an eltra cs 800 carbon sulphur analyser (induction furnace). a series of seven concrete mixes were produced and are summarized in table 1. the batch size was 1 m3. the default (reference) mix design had a w/b of 0.55. the binder content totalled 330 kg/m3 and was comprised of 81% cement and 19% limestone. two batches were produced with untreated wash water comparing concrete excluding and including a retarder as a dispersant to improve slump (marked nt and nta for not treated and not treated plus dispersant admix). likewise, two analogous batches were produced with co2-beneficiated wash water (marked ct and cta for co2-treated and co2-treated plus dispersant admix). lastly, an additional two batches were produced with co2-beneficiated wash water wherein the level of the superplasticizing admixture was increased and either excluded or included the dispersant addition (marked ctp and ctap for co2treated plus plasticizer and co2-treated plus admix plus plasticizer). adjustments were made to the binder loadings of batches containing wash water in order to compensate for the additional solids present in the wash water slurry. the total powder content (cement + limestone + wash water solids) was maintained at 330 kg/m3. reductions were made to the limestone in the first four cases (nt, nta, ct, cta) while a proportional reduction to both cement and limestone was investigated in the final case (ctp, ctap). in the final case the admixture loading was changed with respect to the reference and the adjustments were determined with respect to what was required to achieve the target slump. the dispersant loadings were in proportion to the wash water solids. the concrete was measured in terms of slump and compressive strength at 24 hours, 7 days, 28 days and 378 vol. 33/2022 sustainable concrete production ϭй ϱй ϭϭй ϭϱй ϯϭй ϯϱй ϯϭй ϯϱй ϰϭй ϰϱй ϱ ϲ ϳ θ ϵ ϭϭ ϭϭ ϭϯ ϭϯ ϭϰ ϭ ϯϭ ϲϭ ϵϭ ϭϯϭ ϭϱϭ ϭθϭ �k ϯ� ăŷ ě� , ϯk �đ žŷ ƚğ ŷƚ �ď lj� žƌ ŝő ŝŷ ăů �ŵ ăɛ ɛ� ɛž ůŝě ɛ ɖ, dƌğăƚŵğŷƚ�dŝŵğ ɖ, �kϯ�ǀŝă�d'� �kϯ�ǀŝă�ŝŷěƶđƚŝžŷ �žƶŷě�,ϯk figure 1. carbon uptake in wash water solids and solution ph with treatment time of wash water slurry treated with co2. 56 days with the exception of the batches containing increased plasticizer (pce) which were assessed only at 7 and 28 days. 3. results 3.1. co2 treatment of the slurry the co2 uptake and ph data over time for the two treated slurries closely agreed. one case is presented in figure 1 and discussed herein. the ph of the treated slurry was 12.7 after three hours of prehydration. the ph declined upon addition of the co2 and was only a modest decline at first, down to 12.1 after 60 minutes, before a more rapid drop leading to a ph of 7.1 at 150 minutes. the injection of the co2 into the agitated slurry functionally changed when the treatment neared completion; whereas the initial injection of carbon dioxide did not result in any bubbles of carbon dioxide gas reaching the surface of and exiting from the slurry, the slurry displayed extensive bubbling around the time of neutralization. the carbon content of the slurry solids was observed to increase with treatment time in a broadly linear fashion over the first 120 minutes of treatment. ultimately the treatment was complete after 150 minutes at which time the solids were 28.1% co2 by mass. the carbon contents, as presented in the figure, have been normalized to the same ignited mass to express the uptake as a (percentage) mass of carbon dioxide per mass of the original anhydrous cement. this allowed for a simple comparison against the amount of co2 delivered, which can be described in the same terms. the uptake at the conclusion of the treatment was 41.4% co2 by weight of cement under the tga analysis and 37.4% with the induction furnace. the uptake can be used to calculate process efficiencies. the injection of co2 gas at 600 lph with a density of 1.98 g/l over 2.5 hours means that 2.97 kg of co2 was delivered. where the initial slurry contained 7.7 kg of cement, the uptake rate via the tga translates into 3.19 kg of bound co2 for an efficiency of 107% (a more robust tga analysis that includes unattributed weight losses and a true loi would serve to decrease this calculated value). using the uptake rate via induction furnace there were 2.88 kg of co2 mineralized for a conversion efficiency of 97%. the bound water content was normalized in the same fashion as the bound carbon dioxide. it was observed that the bound water increased over the first 90 minutes of the treatment before declining. in particular the decline in ph may have changed the equilibrium of any hydrate phases. for example, ettringite is destabilized below a ph 10.7 while monosulphate disappears below 11.6 [13]. the change in rate of the uptake curve between 120 and 150 minutes attests to a slowing down of the reaction. the delivery of the carbon dioxide was not changed so at this time the entering carbon dioxide is tending to change the ph of the solution moreso than mineralize with the slurry solids. 3.2. fresh properties the slump measurements of the seven batches of concrete are presented in figure 2. the slump of the control was 19 cm after the conclusion of mixing (5 minutes after the addition of water) and was used as a target in later batches that were adjusted for slump. the untreated wash water served to reduce the slump by about 25% while an addition of the dispersant in turn increased the slump by 10%. the slump of the batch with the co2 treated slurry had 379 sean monkman, vincent meyer acta polytechnica ctu proceedings ϭ ϭ͕ϯ ϭ͕ϰ ϭ͕ϲ ϭ͕θ ϭ ϭ͕ϯ ϭ͕ϰ ϭ ϱ ϭϭ ϭϱ ϯϭ ϯϱ �žŷƚƌžů ed ed� �d �d� �dw �d�w � ěŵ ŝdž �;й ͬď ŝŷ ěğ ƌϳ ^ů ƶŵ ɖ� ;đ ŵ ϳ w�� �ŝɛɖğƌɛăŷƚ figure 2. workability (slump) of concrete batches produced with wash water slurry. admixture loadings of dispersant and pce are indicated as symbols read from the secondary vertical axis. batch control nt nta ct cta ctp ctap 24 h strength (mpa) 10.2 10.8 10.0 14.6 13.4 n/a n/a 7 d strength (mpa) 25.1 26.0 29.9 30.0 31.0 26.4 29.7 28 d strength (mpa) 32.6 32.3 35.8 36.0 36.9 33.0 36.2 56 d strength (mpa) 35.0 n/a 36.9 38.9 39.9 n/a n/a relative strength 24 h 100 % 106 % 98 % 143 % 131 % − − relative strength 7 d 100 % 104 % 119 % 120 % 124 % 105 % 118 % relative strength 28 d 100 % 99 % 110 % 110 % 113 % 101 % 111 % relative strength 56 d 100 % − 105 % 111 % 114 % − − table 2. compressive strength of concrete mixes from 24 hours to 56 days. a major reduction (75%) in workability and, as in the untreated slurry case, an improvement in slump was observed when the dispersant was used. however, in either case, it was observed that the use of the treated slurry resulted in a workability loss. when the addition of a superplasticizer was explored the slump was improved from 9.5 cm to 18.5 cm to effectively match the target. the addition of the dispersant before the suplerplasticizer allowed the same flowability to be achieved albeit with 33% less pce. the workability issues can be solved through the use of conventional admixtures. 3.3. compressive strength the summary of the compressive strength data is presented in table 2. reported compressive strengths are averages of three specimens. the untreated wash water had a small positive effect at the two earliest ages but was neutral at 28 days. the addition of the dispersant resulted in improved performance at ages of 7 days and beyond, albeit diminishing relative to the control from 19% better at 7 days to 5% better at 56 days. the carbonation treatment served to improve the concrete strength at all ages of batches so produced. the benefit was greatest at the earliest age with a 43% benefit at 24 hours. as with the untreated slurry, the relative strength benefit declined with increasing age but remained at 11% at 56 days. the batch that included the dispersant was not quite as strong at 24 hours (a 31% benefit) but showed higher strengths than the batch without the dispersant at all subsequent ages concluding at a 14% benefit at 28 days. while the batches with the treated slurry were stronger than the control they also showed a reduced workability that would render them unviable. the batches with the treated slurry and the pce had a 10% lower cement content than the control. even so, the 7 and 28 day strengths were acceptable and slightly above the control. as with the other examples of using the dispersant, the strength was increased further with an 11% improvement observed at 28 days, despite the cement reduction. 4. discussion the compressive strength results are best compared in terms of the binder efficiency. the comparisons between the control and the batches with acceptable workability and using the treated slurry are outlined in table 3. the cementing efficiency was calculated in terms of the cement clinker (assuming a clinker factor in the cement of 92% as per the cembureau epd 380 vol. 33/2022 sustainable concrete production batch control ctp ctap clinker efficiency 28d (kg clinker m−3 mpa−1) 7.5 6.7 6.1 cement efficiency 28d (kg cement m−3 mpa−1) 8.2 7.3 6.6 binder efficiency 28d (kg binder m−3 mpa−1) 10.1 9.0 8.2 powder efficiency 28d (kg powder m−3 mpa−1) 10.1 10.0 9.1 carbon efficiency (kg co2 m−3 mpa−1) 7.4 6.3 5.7 anticipated strength via cement loading (mpa) 32.6 29.3 29.3 actual strength (mpa) 32.6 33.0 36.2 change in strength (mpa) n/a 3.7 6.9 equivalent units of cement (kg) n/a 30.0 56.2 actual units of ww solids (kg) n/a 33 33 units ww solids/unit cement n/a 0.91 1.71 table 3. cementing efficiency summary of control and batches with treated slurry. [14]), the cement content, the overall binder content (cement + limestone), the overall powder content (cement + limestone + wash water solids) and the carbon impact [15] (relevant to the binder mixture of cement + limestone + wash water solids). the carbon intensity of the cement is taken to be 0.898 t co2/tonne cement [14] while that of the limestone is taken as 0.008 t co2/t cement [16]. it was observed that the clinker, cement and binder efficiencies of the batches produced with wash water were improved 11% over control in the ctp batch. whereas the control produced concrete with a 28 day strength of 32.6 mpa and 267 kg cement/m3 concrete, the ctp batch was 33.0 mpa with 240 kg of cement. if the cement was to provide strength with the same efficiency in the latter as it had in the former then the expected strength was 29.3 mpa. the difference between the actual strength and the expected, +3.7 mpa, can be attributed to the wash water since the overall powder loading was unchanged. whereas 3.7 mpa of strength would be associated with 30 kg/m3 of cement the ctp batch contained 33 kg of wash water solids. the powder efficiency was unchanged in comparison to the control while the carbon the efficiency was improved 15%. there appeared to be a synergy upon adding both the dispersant and the pce (batch ctap). all of the clinker, cement, and binder efficiencies were improved 19% as compared to the control. in this instance the actual strength was +6.9 mpa greater than the expected strength as per the cement efficiency of the control and cement loading of the batch. the strength difference is equivalent to 56.2 kg of cement. the incorporation of 33 kg of wash water solids meant that the solids had a cement replacement efficiency of 1.7 units of cement per unit of solids. the powder efficiency was improved 10% in comparison to the control while the carbon the efficiency was improved 22%. the sustainability impacts can be achieved with respect to the mineralized carbon dioxide in the wash water solids, the avoided cement as part of the mix design modification, and recycling of waste cement and avoided potable mix water. the direct co2 impacts are estimated from the uptake of carbon dioxide into the wash water. the uptake of 28% co2 by weight of solids determined that the 33 kg of wash water solids/m3 concrete contained 9.2 kg of mineralized co2. in accordance with the cembureau epd for portland cement, the 27 kg of cement avoided were associated with 24.2 kg of carbon dioxide. the combined carbon reduction was 35.4 kg of co2 of which the avoided was 75% and the mineralized was 25%. the carbon impact of the binder was reduced 14% from 240.3 kg co2/m3 concrete in the reference to 206.8 kg co2 m3 concrete in the two batches with co2 treated slurry. the replacement of the fresh mix water meant that 182 kg fresh water/m3 concrete was saved. the untreated wash water produced acceptable concrete with workability better than in the co2 treated case and acceptable strength albeit not as high as in the co2 treated case. the performance of the concrete produced with the untreated wash water would be expected to vary with age of the wash water. it would be less predictable and burdensome to the producer given the evolving properties and condition of the wash water slurry itself. on the other hand, the treatment of the slurry with co2 not only mineralizes co2 permanently into waste cement, but it stabilizes the wash water such that its performance is consistent and predictable. thermal analysis data characterizing % bound co2 and % bound water showed that for an untreated slurry the total mass loss (loi) was increased from 1.2% in the anhydrous case to 3.9% for the slurry hydrated 3 hours to 22.6% for the slurry hydrated 26 hours. conversely, a test of slurry solids treated to an uptake of 27.4% by mass co2 and 8.9% by mass bound water (an loi of 36.3%) at the conclusion of the treatment was observed to be 27.9% co2 and 7.7% h2o (loi of 35.6%) after holding for an additional 24 hours. a successful adoption of the approach could realize several operational benefits. the operating expense for solids disposal could be greatly reduced or eliminated. both costs related to neutralization of 381 sean monkman, vincent meyer acta polytechnica ctu proceedings wash water prior to discharging and permitting costs associated with discharging could be avoided. the process would come with costs associated with the merchant carbon dioxide and increased admixture usage but these costs could be offset by savings of any admixtures used in an incumbent wash water management approach and through cement savings. the future may bring monetizable carbon offsets for co2 utilization that would improve the economic motivation to implement the approach. 5. conclusions the use of carbon dioxide to treat concrete wash water allowed the treated slurry to be used as mix water in the creation of performance enhanced concrete. a high specific gravity wash water slurry was shown to bind co2 at around 40% by weight of cement. the complete replacement of potable water with treated slurry lead to concrete with a reduced workability, but modification of the admixture loading addressed the issue. a mix design adjustment to reduce the binder loading in direct proportion to the suspended solids contained within the slurry did not compromise the compressive strength performance. the wash water solids contributed to the performance of the binder and thereby displayed a latent cementitious property. the approach aligns with industry sustainability goals to reduce the clinker content in concrete mixes, reduced wastes generated by concrete plants, to increase recycled materials usage and to reduce the water impacts of concrete production. an environmental benefit is realized through both mineralized co2 and avoided co2 from replaced cement and amounted to a 14% reduction of the overall binder carbon impacts. the proposed approach represents a significant operational, resource efficiency, cost saving, regulatory, and logistical opportunity for sustainable concrete production. acknowledgements the authors acknowledge that the work was completed within the first edition of the lafargeholcim accelerator program. technical work was completed at the lafarge centre de recherche (lcr) in lyon, france. the contributions of sandra boivin, sébastien rols, yves francomme, vincent briaud, and julien nawrot are greatly appreciated. references [1] iea and wbcsd. technology roadmap: low-carbon transition in the cement industry, paris and geneva: oecd/iea and wbcsd, 2018. [2] a. favier, k. scrivener, g. habert. decarbonizing the cement and concrete sector: integration of the full value chain to reach net zero emissions in europe. iop conference series: earth and environmental science 225, 2019. https://doi.org/10.1088/1755-1315/225/1/012009. [3] c. hepburn, e. adlen, j. beddington, et al. the technological and economic prospects for co2 utilization and removal. nature 575(7781):87-97, 2019. https://doi.org/10.1038/s41586-019-1681-6. [4] w. ashraf. carbonation of cement-based materials: challenges and opportunities. construction and building materials 120:558-70, 2016. https: //doi.org/10.1016/j.conbuildmat.2016.05.080. [5] d. zhang, z. ghouleh, y. shao. review on carbonation curing of cement-based materials. journal of co2 utilization 21:119-31, 2017. https://doi.org/10.1016/j.jcou.2017.07.003. [6] co2 sciences the global co2 initiative. global roadmap for implementing co2 utilization. ann arbor, mi: university of michigan. available online at: https://deepblue.lib.umich.edu/bitstream/handl e/2027.42/150624/co2u_roadmap_final_2016_12_07 [7] c. lobo, g. mullings. recycled water in ready mixed concrete operations. concrete infocus 2(1):17-26, 2003. [8] astm c1602m-12. standard specification for mixing water used in the production of hydraulic cement concrete, 2012. [9] rmc research and education foundation. sustainable concrete plant guidelines version 1.1, 2011. https://rmc-foundation.org/wp-content/upl oads/2017/07/scp-guidelines-version-1.1.pdf. [10] r. l. berger, j. f. young, k. leung. acceleration of hydration of calcium silicates by carbon dioxide treatment. nature physical science 240(97):16-8, 1972. https://doi.org/10.1038/physci240016a0. [11] c. j. goodbrake, j. f. young, r. l. berger. reaction of beta-dicalcium silicate and tricalcium silicate with carbon dioxide and water vapor. journal of the american ceramic society 62(3-4):168-71, 1979. https: //doi.org/10.1111/j.1151-2916.1979.tb19046.x. [12] d. r. moorehead. cementation by the carbonation of hydrated lime. cement and concrete research 16(5):700-8, 1986. https://doi.org/10.1016/0008-8846(86)90044-x. [13] a. gabrisová, j. havlica, s. sahu. stability of calcium sulphoaluminate hydrates in water solutions with various ph values. cement and concrete research 21(6):1023-7, 1991. https://doi.org/10.1016/0008-8846(91)90062-m. [14] cembureau 2015 environmental product declaration portland cement (cem i) produced in europe. [15] b. l. damineli, f. m. kemeid, p. s. aguiar, et al. measuring the eco-efficiency of cement use. cement and concrete composites 32(8):555-62, 2010. https: //doi.org/10.1016/j.cemconcomp.2010.07.009. [16] s. a. miller, v. m. john, s. a. pacca, et al. carbon dioxide reduction potential in the global cement industry by 2050. cement and concrete research 114:115-24, 2018. https: //doi.org/10.1016/j.cemconres.2017.08.026. 382 https://doi.org/10.1088/1755-1315/225/1/012009 https://doi.org/10.1038/s41586-019-1681-6 https://doi.org/10.1016/j.conbuildmat.2016.05.080 https://doi.org/10.1016/j.jcou.2017.07.003 https://rmc-foundation.org/wp-content/uploads/2017/07/scp-guidelines-version-1.1.pdf https://doi.org/10.1038/physci240016a0 https://doi.org/10.1111/j.1151-2916.1979.tb19046.x https://doi.org/10.1016/0008-8846(86)90044-x https://doi.org/10.1016/0008-8846(91)90062-m https://doi.org/10.1016/j.cemconcomp.2010.07.009 https://doi.org/10.1016/j.cemconres.2017.08.026 https://doi.org/10.14311/app.2022.33.0105 acta polytechnica ctu proceedings 33:105–111, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague rc deck stiffened arch existing bridges: simulated design and structural analysis giovanni criscia, ∗, francesca ceronia, gian piero lignolab, andrea protab a university of naples parthenope, department of engineering, centro direzionale c4, 80143 naples, italy b university of naples federico ii, department of structure for engineering and architecture, via claudio 21, 80125 napoli, italy ∗ corresponding author: giovanni.crisci@uniparthenope.it abstract. the 20th century is known as the age that gave birth to the largest reinforced concrete structures. many applications of this new material were realized at that time, both from a theoretical and practical point of view. with reference to bridges, the engineer robert maillart achieved a new concept of arched bridges, characterized by very stiff deck beams and slender and wide vaults, i.e., the "deck-stiffened arch". the paper deals with the study of such bridge typology, particularly widespread in italy around the 50s of the 20th century. while, nowadays, calculation tools allow developing very refined structural modelling, in the past very simple structural schemes were adopted in the design phase in order to simplify the calculation effort. the study starts from a "simulated design" of such a bridge typology adopting a reliable geometry and following the design rules and the simplified structural schemes of the time and, then, by means of a refined three-dimensional model, the performance of a typical "maillart-type arch" bridge is analysed. keywords: maillart-type arch bridges, reinforced concrete arch bridges, structural analysis. 1. introduction the arch bridges are one of the oldest type of structures and are still today widely used to cross deep ravines in mountain areas. reinforced concrete (rc) arch bridges represent a suitable alternative to the classic "girder bridges", both on the aesthetical and functional points of view. the arch is, indeed, able to optimize the compressive strength of concrete or masonry materials by transferring the gravitational loads through a distribution of axial compressive forces. due to the development of new building techniques and high performance materials, the diffusion of rc arch bridges increased considerably. rc arch bridges can have different forms such as "tied arch bridges", "deck arch bridges" and "through-arch bridges". generally, the "deck arch bridge" is the most widely preferred typology because the arch sustains the gravity loads by means of compressive stresses. the key components of an rc deck arch bridge are: the deck beam (i.e., the superstructure), the piers (i.e., the piers that connect the deck to the arch), and the arch. within the "deck arch bridge" typology, the paper focuses the attention on the "deck-stiffened arch bridges", also known as "maillart-type arch bridges", aăparticularly widespread typology in italy, built between the 40s and the 60s of 20th century. in large span bridges, the contribution of the deck beam to the bearing capacity of the whole structure is completely ignored, but it is assumed that the deck beams only bear and transfer gravitational and accidental loads over the arch. on the other hand, in medium span bridges, the contribution of the deck beam, as long as it is continuous, is considerable and it would be uneconomical not to take its contribution into account. the "maillart-type arch" bridges (see figure 1a and 1b) are an example of the so-called collaborative deck structures, in which the deck is not only a simple element for transferring the loads, but it contributes to the bearing capacity of the whole structure. for typical rc arch bridges, the inertia of the deck beam is very small in comparison with the inertia of the arch cross section and, thus, the arch bears most bending moments (and shear). the swiss engineer robert maillart, expert of bridges and rc structures, at the beginning of 20th century, reversed this idea, since he conceived a new bridge typology, which took his name, made of a slender and wide vault, characterized by a low moment of inertia, and a very stiff deck beam. the two elements were connected by means of slender piers, often wall-like and lightened by central windows. the inversion of the classic ratio of arch and deck stiffness was a very important expedient, since it led to the significant reduction of the bending (and consequently shear) forces in the arch, which worked mainly in compression (i.e. a funicular arch). this has the advantage of simplifying the scaffoldings for the construction of the arch, which was usually built in deep ravines and ended with a structural scheme that is an inverted suspended bridge. 105 https://doi.org/10.14311/app.2022.33.0105 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en g. crisci, f. ceroni, g. p. lignola, a. prota acta polytechnica ctu proceedings � d�� � e�� figure 1. example of maillart-type arch bridge. several viaducts operating in italy and inserted in the national highway or motorway network have now reached the so-called "service life". surely these structures have been designed and built with a very different philosophy from the current design principles and with different design actions. recent collapses, such as the one involving the polcevera viaduct in italy, have highlighted the need of a good balance between social, economic, and environmental evaluations in order to attain a sustainable development. for bridges, sustainability can be interpreted as the resilience that the structure has to offer against the ever increasing environmental and loading conditions. thus, for a suitable sustainable approach, it is necessary to understand the structural concept that governed the design of the bridges in order to catch their resilience over the time. within the rc deck arch bridge, in this paper, the attention is focussed on the "maillart-type arch" typology. firstly, the design methodology of the time for such a bridge typology was investigated, then a "simulated" design of a reliable geometry was carried out according to simplified models and assumptions in agreement with the methods of the time. the elements of the simulated bridge have been designed and verified according to the "allowable stress method (asm)" or "working stress method (wsm)" [1] and under the design loads of the time. lacking high computational capacity, the typically adopted design schemes at the time were easier than those currently adopted. two-dimensional models have been, indeed, adopted for the design of structural elements and some simplifications have been assumed about their constraints and working conditions. in turn, the real geometry of the case study has been implemented in a finite element (fe) software and, under the loading conditions of the time, the structural elements have been verified again according to the wsm. 2. design of structural elements for the case study 2.1. definition of geometry, materials and loading conditions a simple model of a "maillart-type arch" bridge, i.e. a rc deck arch bridge, was assumed according to the scheme of figure 2a as case study for the analysis presented in this paper. the bridge is made of a slender vault with a cross section having constant dimensions of 6m × 0.2ăm. the shape of the arch was designed in order to be funicular of the gravitational loads corresponding to the self-weight of the entire structure. this means that it is subjected to axial compression only, without any bending moment and shear force. to this aim, an iterative procedure was adopted where, taking into account the gravitational loads of the deck and the piers, a funicular arch was generated starting from the shape of a parabolic arch [2]. this was a common practice in the past, mainly in order to facilitate the construction of the bridge; if the vault was built first and, later, the entire superstructure was built on it, it is possible to use very light scaffolding for the vault during the hardening of concrete. in fact, after this phase, the arch alone is able to bear the loads for the next phases of the construction. the length of the bridge, at the imposts of the arch, is l = 60 m, while the rise is f = 24 m; the deck is 9 m wide (figure 2b) and it consists of one driveway with two lanes (3.5 m each) and with two lateral sidewalks 1 m wide. the structure of the deck is made of a regular grid of main and secondary rc beams (figure 2b) with rectangular cross sections with dimensions of 0.3 m × 1.2 m and 0.3 m × 0.6 m , respectively, and of a slab 0.2 m thick. a total of 11 triplets of piers with a constant 0.3 m × 0.5 m cross section connects the vault to the deck (see figure 2a). typical concrete and steel usually used for this type of structure in ’50s of 20th century were assumed. based on historical sources and on several consulted projects of the time, tor 60 steel was largely used: it was characterized by an average yield 106 vol. 33/2022 rc deck stiffened arch existing bridges d�� e�� /� ����p i� �� �� p ��p ��p ��p ��p ��p ��p ��p ��p ��p ��p 3� 3� 3� 3� 3� 3� 3� 3� 3� 3�� 3�� d �p �p �p ���p / 0dlq�ehdpv 6hfrqgdu\�ehdpv figure 2. a) longitudinal view of the deck arch bridge assumed as case study; b) plan view of the deck. point of about 600 mpa and was classified as semihard or hard steel. a characteristic cylindrical compressive strength fck = 30 mpa was adopted (i.e. about an average cubic strength rcm = 45.8 mpa), since semi-hard steel was usually adopted with more performant concretes. the structural elements of the bridge were designed replicating the common practice of the past. at that time, very simple structural schemes were adopted in comparison with current design approaches. for the purposes of this work, a simulated design was performed only for the beams, the piers and the arch. based on the design rules of that time and after the examination of old projects of rc deck arch bridges too [3], the typical design procedure was replicated in order to define the dimensions of elements and amount of steel reinforcement. after the category of the bridge is fixed depending on the road typology, the expected traffic load was defined. thus, the analysis of load combinations was carried out with reference to the mandatory codes of the time, in particular the royal decree of 1939 [4] and the circular n. 6018 on the loads on bridge structures of 1945 [5]. in particular, assuming a bridge of first category (high traffic roads), two different types of loading schemes were adopted: 1. two or more undefined trains of trucks weighing 12 tons (about 120 kn) side by side; 2. two compressor rollers weighing 18 tons each (about 180 kn) side by side. in addition, a compact crowd was expected on the sidewalks next to the driveway. the second load scheme was typically considered for local verifications, since it locally provides higher stresses (a concentrated force of 6 tons per wheel was, indeed, assumed). conversely, the design of the main structural elements i.e. piers, arch, and beams was carried out with reference to the first load scheme, which globally engages the entire structure. it is worth to note that in the following, the simulated design will be carried out according to a simplified model and will be, thus, focussed only on the beams, piers and arch as load carrying elements, while the slab of the deck will be considered only with its weight. 2.2. assessment of loading schemes for the main beams there are many simple structural schemes that can be used for a preliminary design. the approach adopted in this paper considers an equivalent two-dimensional structure, characterized by a continuous deck beam supported by the piers, a fully fixed arch at its imposts and axially rigid piers connected to the deck and the arch with hinges at both ends, as commonly adopted by the designers of that time. such an assumption yields to zero shear and bending actions on the piers, while the real constraints and the detailing allows for the development of some shear and bending actions; however usually this mismatch has negligible impact on the structural behaviour. moreover, in this phase, the main beams are considered rectangular instead that t shaped, hence neglecting the structural contribution of the slab, but with the height equal to the whole deck thickness. for the main beams, the assessment of maximum stress was carried out by placing the moving loads, given by the loading condition 1 [4, 5], in the most 107 g. crisci, f. ceroni, g. p. lignola, a. prota acta polytechnica ctu proceedings figure 3. influence line of bending moment and position of the moving loads able to maximize the positive bending moment. unfavourable position on the deck, which was found by means of the influence lines. figure 3 shows the influence lines of the bending moment for moving vertical loads in some cross sections of the deck (red lines). the envelope of the peaks (blue line in figure 3) provides the diagram of the maximum and minimum moments from which it is possible to identify the cross section and, therefore, the position of the moving loads able to maximize the bending moment. it was found that the most stressed section is at about 1/4 of the bridge span and the loads associated to its development should be applied on about one third of the bridge span. given the maximum moment (positive and negative), it has to be divided among the individual main beams of the deck. the division can follow different approaches: for the examined case, it was assumed that the load solidly affects all the elements of the deck so that each beam is burdened by the same percentage of load. it is worth to note that the maximum bending moment induced by the moving loads must be combined with the actions produced by the gravitational loads; to take into account the dynamic effects, the moving loads have been amplified by 35%, which is obtained by the following formulation: ϕ = 1 + 16 l + 40 (1) where l is the span of the main beams to be designed. with regard to the royal decree of 1939 [4], the design of steel reinforcement was made according to the "allowable stress method (asm)", also known in literature as "working stress method (wsm)" [1]. for the main beams, the steel bars have been designed with the simplified formula for rc cross section with single bars layer: as = m 0.9 d σy, k (2) where d is the effective height and σy, k is the steel working stress, which is established by means of a safety factor 2 with respect to the characteristic yielding stress and, thus, is 300 mpa, being σy, k = 600 mpa. moreover, for the main beams, a double steel bars layer has been placed, a common solution found in several old projects [3]. details of steel reinforcement for the main beams are reported in figure 5. 2.3. assessment of loading schemes for the arch, the piers and secondary beams the position of the moving loads giving the maximum bending moment on the deck does not yield the maximum horizontal thrust in the arch too. in fact, in order to maximize the horizontal thrust, h, and the compression force, n , in the arch as well, the moving loads must be applied on the whole deck. since the arch is designed to be funicular of self-weights and even under moving loads it remains funicular due to the high flexural stiffness of the deck (because bending moments in the arch remain negligible), the compression force in the arch can be calculated by knowing the value of the horizontal thrust h and inclination of cross section, α: n = h cos α (3) the piers were studied under the same load condition considered for the arch. it is worth to note that both the piers and the arch, under the simplified design assumption, were designed to carry compression force only. under such an assumption, the area of the steel bars was designed as 0.8% of the ideal concrete 108 vol. 33/2022 rc deck stiffened arch existing bridges / 0rylqj�/rdg 'hdg�/rdg'hdg�/rdg / 0rylqj�/rdg figure 4. different constraint conditions for the secondary beams. ��i���#�p ��i�� �i�� ����fp �� �� fp 6hfrqgdu\�ehdp ��i�� ��i�� ����fp �� �� fp 3lhu ��i�� ��i�� ����fp �� �� �f p ��i��0dlq�ehdp �� �� fp ����fp $ufk�9dxow ��i���#�p figure 5. details of cross section and steel reinforcement of different structural elements. area, i.e. the minimum area of concrete related to the maximum compressive force, n [4]: aci = n σc, c (4) where σc, c is the working stress for the concrete under compression which was assumed equal to 70% of σc, f . since the compressive cylinder strength of concrete was assumed as fck = 30 mpa (i.e. characteristic cubic compressive strength rck = 36 mpa), the working stresses under flexural action, σc, f , is: σc, f = 7.5 + rck − 22.5 9 = 9 mpa (5) and the maximum working stress in the concrete under compression forces is, thus, σc, c = 6.3 mpa. the secondary beams were considered appropriately fixed to the main beam. in particular, the constraint conditions at the ends have been assumed as ranging between a simple supported beam and a fully fixed beam, in order to maximize the positive and negative bending moments along the element (figure 4a). the secondary beams have been designed with reference to their tributary area to define the entity of the gravitational loads (uniformly distributed loads) and with reference to the worst condition for the moving loads, i.e. when the axis of the vehicle is just on the secondary beam (loading condition 2 according to [4, 5]). the steel reinforcement was designed according to eq. 2 too. figure 5 reports the details of cross sections and steel reinforcement for the arch, the piers and the beams. 3. analyses and structural verifications based on the bridge geometry discussed in the previous section, a three-dimensional model has been implemented in the sap2000 software [6]; two-node frame elements are used to model the beams of the deck and the piers of the bridge. the eccentricity of the piers with respect to the axis of the arch is taken into account with rigid offset and massless links. differently from the simplified schemes adopted in the simulated design, the piers are fixed at the ends and are not connected, thus, by hinges. figure 6. 3d fe model of the rc deck arch bridge assumed as case study. 109 g. crisci, f. ceroni, g. p. lignola, a. prota acta polytechnica ctu proceedings d�� e�� ��� ��� ��� ��� ��� : ru nl qj �6 wu hv v� fr qf uh wh �>0 3d @ 3lhuv $ufk 7kuhvkrog ydoxh�����03d � �� �� �� �� �� : ru nl qj �6 wu hv v� >0 3d @ 3lhuv $ufk 7kuhvkrog ydoxh�����03d figure 7. structural verifications of compression stresses in the arch and the piers: a) compressed concrete; b)ăcompressed steel reinforcement. d�� e�� � � � � � �� �� : ru nl qj �6 wu hv v� fr qf uh wh �>0 3d @ 3rv��0rp��0dlq�ehdpv 1hj��0rp��0dlq�ehdpv 3rv��0rp��6hf��ehdpv 1hj��0rp��6hf��%hdpv7kuhvkrog ydoxh���03d � �� ��� ��� ��� ��� : ru nl qj �6 wu hv v� >0 3d @ 3rv��0rp��0dlq�ehdpv 1hj��0rp��0dlq�ehdpv 3rv��0rp��6hf��ehdpv 1hj��0rp��6hf��%hdpv 7kuhvkrog ydoxh�����03d figure 8. structural verifications for bending moment in the beams: a) compressed concrete, b) steel reinforcement in tension. based on the fe model shown in figure 6, the simple rc deck arch bridge assumed as case study was analysed under the loading conditions prescribed by circular n. 6018 of 1945 [5]. in order to carry out a global analysis of the structure, the undefined trains of trucks weighing 12 tons were considered, i.e. the loading condition 1 provided by [5]. such a loading scheme was converted into a uniform distributed load by means of appropriate tables [5]. each main beam, in addition to its own weight and the loads relative to the pavement, was loaded by the same amount of moving loads, according to a common assumption particularly widespread in ’40s and ’50s years. the compact crowd, on the other hand, was considered only for the two main edge beams. the structural verifications, as done for the simulated design, have been performed according to the wsm [1]. in addition to the allowable normal stresses under bending actions for the concrete given by eq. 5, the allowable shear stresses, for a concrete with a good performance, were equal to: τc, 0 = 0.6 mpa, τc, 1 = 1.6 mpa (6) note that if τ < τc, 0, it is not necessary to design an adequate shear reinforcement and usually 3 or 4 stirrups per meter were adopted. if τc, 0 < τ < τc, 1 an adequate shear reinforcement is required; otherwise, it is necessary to change the rc cross section. for the steel, the same allowable stress of 300 mpa has been considered. according to the wsm, the structural verifications are satisfied when the stresses in all the cross sections of the element are lower than the allowable thresholds, assuming a linear elastic behaviour for materials and a homogenization coefficient n = 10 [4]. figure 7 summarizes the working stresses in compression in the concrete and in the steel reinforcement distributed along the horizontal axis for the cross sections of the arch and the piers under compression forces. the working stresses are significantly lower than the threshold values (6.3 mpa for concrete and 300 mpa for steel). moreover, the working stresses in the beams, both main and secondary ones freely distributed along the horizontal axis, under bending moment (positive and negative) and shear forces are shown in figure 8 and 9, respectively. figure 8 highlights that the stresses in every cross section due to bending moments are always lower than the allowable ones (9 mpa for concrete and 300 mpa for steel). moreover, figure 9 shows that in every cross section the shear stresses are always lower than τc, 0, meaning that it was not 110 vol. 33/2022 rc deck stiffened arch existing bridges ���� ���� ���� ���� ���� : ru nl qj �6 kh du �6 wu hv v� >0 3d @ 0dlq�ehdpv 6hfrqgdu\�ehdpv7kuhvkrog ydoxh�wc0 ����03d figure 9. structural verifications for shear in the beams.. necessary [4] to design a specific shear reinforcement, but it was sufficient to consider a minimum amount, i.e. 3 or 4 stirrups per meter. it is worth noting that the shear verification was carried out only for the beams since the arch, designed as funicular of the self-weight of the whole structure, and the piers, designed as pendulums, are expected to work mainly in compression. the fe model in sap2000 has, indeed, highlighted the presence of bending moments and shear actions on the arch as well, although of very small entity if compared to the dominant axial compression load. although starting from a very simple structural scheme, it seems that the actions on the structural elements calculated with a more refined fe model are consistent with the simplified approaches used in the past. it is worth to remember that the wsm entails the designer to check that the maximum working stresses in the structural elements under the service loads do not exceed the allowable stress. therefore, although some values are close to the reference thresholds, it can be reasonably supposed that they are however far from the effective failure values. future analyses will be, indeed, focussed on the verification of the same case study under the actions provided by current italian building code and adopting the current methodology, i.e. the limit states approach. 4. conclusions several rc deck arch bridges were built around the 1950s in italy. the present work analyses the simulated design of a rc "maillart" arch bridge, in particular, replicating a typical approach adopted in the past. starting from simple design schemes, a simulated design was performed accounting for the load conditions required by the past codes. the response of the case study was, then, examined by a 3d finite element model under the same load conditions adopted in the simulated design phase, where real shape and constraints have been simulated. the structure was verified according to the working stress method (wsm), typical of the 1950s. all the verifications resulted adequate. further studies will investigate several modelling strategies and will concern the analysis of the case study under the actions provided by new italian building code and performing the verifications according to the current verification methodology, i.e. the limit states approach. references [1] w. lin, t. yoda. bridge engineering classifications design loading and analysis methods, butterworth heinemann, 2017. :https: //doi.org/10.1016/b978-0-12-804432-2.00001-3. [2] v. franciosi. il ponte ad arco (liguori editore), 1994. [3] w. d. ponti. italiani in cemento armato. santarella and miozzi (book review). town planning review 15(3), 1933. https: //doi.org/10.3828/tpr.15.3.hv41614521888l41. [4] regio decreto 16/11/1939 n. 2229 norme per la esecuzione delle opere in conglomerato cementizio semplice ed armato (in italian), 1939. https://www.edilportale.com/normativa/regio-dec reto/1939/2229/norme-per-la-esecuzione-delle-o pere-in-conglomerato-cementizio-semplice-od-a rmato_15060.html. [5] ministero dei lavori pubblici: direzione generale dei servizi tecnici. circolare 9/06/1945 no. 6018 carichi da considerare nel calcolo di ponti per strade ordinarie (ministero dei lavori pubblici direzione generale dei servizi tecnici) (in italian), 1945. [6] computer & structures inc. (csi) sap2000 integrated software for structural analysis and design computers and structures inc., berkeley, california, 2013. https://www.scirp.org/(s(351jmbntvnsjt1aad kposzje))/reference/referencespapers.aspx?refe renceid=961255. 111 https://doi.org/10.1016/b978-0-12-804432-2.00001-3 https://doi.org/10.3828/tpr.15.3.hv41614521888l41 https://www.edilportale.com/normativa/regio-decreto/1939/2229/norme-per-la-esecuzione-delle-opere-in-conglomerato-cementizio-semplice-od-armato_15060.html https://www.scirp.org/(s(351jmbntvnsjt1aadkposzje))/reference/referencespapers.aspx?referenceid=961255 acta polytechnica doi:10.14311/app.2020.27.0107 acta polytechnica 27(0):107–111, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app estimation of anisotropic elastic properties of carbon fibers using nanoindentation thimmappa shetty guruprasad∗, vincent keryvin, alain bourmaud university of bretagne sud, umr cnrs 6027, irdl, 27 rue armand guillemot bp 92116, 56321 lorient, france ∗ corresponding author: thimmappa-shetty.guruprasad@univ-ubs.fr abstract. understanding the mechanical behavior of carbon fiber reinforced polymers requires knowledge on the deformation behavior of carbon fibers, they are highly anisotropic and heterogeneous. nanoindentation is an efficient method for determining the mechanical properties in small volumes of materials. for isotropic materials, a single nanoindentation test can evaluate an elastic properties of the material. but for anisotropic material, the difficulty increases since measured indentation modulus depends on five elastic parameters (el,et,glt, νlt,and νtt) of the material. nanoindentation experiments are performed on carbon fibers orientated between 0◦ to 90◦ at ten different orientations to the fiber axis. from theoretical models given by vlassak et al. and delafargue and ulm, the elastic constants are predicted numerically by comparing the results of indentation modulus versus orientation angle with the experiments. keywords: anisotropy, carbon fibers, nanoindentation. 1. introduction carbon fiber composites are an important class of materials, which have high tensile strength, high modulus, and lower weight. due to its advantages, the composites are used in many structural applications such as in the aerospace, automotive and naval industries. for the better design of composites having good structural strength, knowing the elastic properties of individual carbon fibers and also the knowledge of predicting its mechanical response is essential. the carbon fibers are generally categorized into four different types based on its strength, namely intermediate modulus (im), high modulus (hm), ultra-high modulus (uhm) and high resistance (hr). the mechanical strength of the composite material reinforced with unidirectional carbon fibers behave to be highly anisotropic in its fiber direction and the strength in the transverse direction is much weaker as compared to longitudinal direction [1]. the most notable experimental technique to determine the mechanical properties of small and delicate structures is through nanoindentation. the nanoindentation is generally used for measuring mechanical properties such as elastic modulus and hardness. since the carbon fibers are highly anisotropic, determining the modulus in any one direction will not determine the complete set of elastic properties of the material. assuming carbon fibers as transversely isotropic materials and due to symmetry we need to determine five elastic constants. so, the indentation modulus is a function of five elastic parameters of the material which are three modulus (el,et,glt) and two poisson’s ratios (νlt, νtt). due to less number of available theoretical models and also due to difficulty in implementing them numerically, determining elastic constants is a challenging task. similar to carbon fibers, there exists an anisotropy in natural fibers, for e.g. jäger et al. [2] determined the anisotropy in the elastic properties of wood cell walls using the nanoindentation technique. they predicted the elastic constants from the numerical solutions of vlassak et al. [3]. recently, csanádi et al. [4] studied the anisotropy in the carbon fibers from a ceramic based carbon fiber composites, here the elastic constants were predicted using the same numerical solutions of vlassak et al., but the experiments were done only on two independent directions i.e. at longitudinal and transverse directions. more recently, duan et al. [5] determined the transverse and the shear modulus of three different types of carbon fibers. the fiber surfaces are prepared by focused ion beam technique and both afm as well nanoidentation tests are performed to determine the indentation modulus. the results from the afm are considered to determine the elastic constants using swanson’s model [6]. in this work, we will determine the elastic constants of carbon fibers of type im2c using nanoindentation, they are pan-based carbon fibers. first, nanoindentation experiments are performed to determine the anisotropic response by cutting the carbon fibers at varying orientation angles concerning to the axis of the fiber. the fibers are cut between 0◦ (longitudinal direction) to 90◦ (transverse direction) in steps of 10◦. the conventional machining techniques are used to obtain the desired surface for the tests. the indentation modulus is calculated using a standard oliver-pharr method. then, we determined the elastic constants numerically using theoretical models; one of them was proposed by vlassak et al. [3]. the model relates the elastic constants with the indentation modulus for the 107 http://dx.doi.org/10.14311/app.2020.27.0107 http://ojs.cvut.cz/ojs/index.php/app t.s. guruprasad, v. keryvin, a. bourmaud acta polytechnica anisotropic materials. the other theoretical model is given by delafargue and ulm [7], which can predict the values only at 0◦ and 90◦. 2. experiments 2.1. materials the epoxy-based carbon fiber composite material is chosen to investigate the elastic properties. the composite consists of polyacrylonitrile (pan) type carbon fibers type im2c (hexcel corporation), and it belongs to the category of intermediate modulus. the average diameter of the carbon fibers are measured to be 5.2 µm. to perform nanoindentation, samples are prepared by cutting the composite at orientation angles between 0◦(longitudinal direction) to 90◦ (transverse direction) in steps of 10◦ with respect to the axis of the carbon fibers. then the composite samples are hot mounted in epoxy with orientation angle parallel to the cutting plane. the desired surface of the carbon fibers for the nanoindentation test is achieved by polishing the samples initially with sandpapers of 1000, 2000 and 4000 grit and then with diamond suspension having a particle size of 1 µm and finally with colloidal silica particles of size 30 nm. 2.2. microscopic analysis as carbon fibers are oriented between 0◦ and 90◦ in steps of 10◦ concerning to the carbon fiber axis, the actual orientation may be different than the cutting angle due to the use of conventional cutting and polishing methods. to measure the actual orientation angle (θ), samples are analysed using the scanning electron microscopy (sem, jeol ltd.), the angle is determined by measuring the size and shape of the carbon fibers from the sem images. for 0◦, the shape is a circle, for 90◦ it is a rectangle as it is perpendicular to the axis and for rest of the angles it is a ellipse. figure 1(a) and (b) shows the sem images of the carbon fibers with orientation angles of 0◦ and 70◦ respectively. to measure the actual orientation angle of the carbon fiber, imagej software is used, where the sem images of the carbon fibers are digitized and the dimensions of the cross section are measured. for 0◦ − 80◦, the major and minor axes of the ellipse is measured from their respective sem images, then the actual orientation angle is determined using the relation θ = cos−1( b a ), where a and b are the major and minor axes of the ellipse. figure 1(c) shows the schematic of the ellipse and the method for calculating the actual orientation angle from the carbon fibers cross section. 2.3. nanoindentation nanoindentation is an efficient method for determining the local mechanical response in small volume of materials. to determine the anisotopic elastic properties of the carbon fibers, indentations are performed perpendicular to its cross section which are having a 10 µm (a) 20 µm (b) a b actual angle of tilt, (c) figure 1. sem images of carbon fibers with cutting angle (a) 0◦ and (b) 70◦. a schematic with dimensions of ellipse and the formula to calculate the actual angle of tilt. orientation angle varying from 0◦ to 90◦ with respect to the axis of the carbon fibers. figure 2 shows the schematic of the indentation performed at an orientation angle (θ) with respect to the axis of the carbon fiber. the nanoindentation experiments are conducted in a continuous stiffness measurement mode using an mts nano xp nanoindenter having a diamond berkovich tip. all the indentations are performed for depths less than 150 nm, the larger depths the may affect the surrounding matrix along with the deformation of the carbon fibers. the hardness and elastic modulus are calculated using a standard method given by oliver and pharr (see 8 for detailed procedure for calculating the modulus and hardness). the effective 108 vol. 27/2020 anisotropy indenter carbon fiber figure 2. a schematic of the carbon fiber indentation at an orientation angle (θ) with respect to the carbon fiber axis. indentation modulus is given by 1 meff = 1 −ν2i ei + 1 m (1) where m = es1−ν2s , which is indentation modulus of the material. the ei and es are the elastic modulus of the indenter and the sample respectively, similarly νi and νs are the respective poisson ratios. the elastic modulus is calculated using the above relation, and the value of ν has a minor influence on the value of the elastic modulus [9]. to determine the anisotropy in the carbon fibers, the samples prepared with varying orientation angle between 0◦ to 90◦ with respect to the axis of carbon fiber are indented. for each orientation angle, at least 20 are indents are performed to ensure the repeatability in the results. 3. theoretical models for anisotropic nanoindentation with the aim of determining the elastic properties of the material by nanoindentation, the indentation modulus for isotropic materials is given by m = e 1 −ν2 , (2) where e is the elastic modulus of the material. for isotropic materials, m is independent of the indentation direction. in the case of anisotropic materials, the relation of indentation modulus with the elastic constants is very complicated as it depends on the indentation direction and the indenter shape. the closed form expressions between the indentation modulus and the elastic material constants (cij) for the transversely isotropic materials are proposed by delafargue and ulm [7] and the solution can be used when the indentation direction is aligned with direction of one of the principal material axes (see 7 for more details). the solution of the indentation modulus is available only at two independent directions i.e. at longitudinal direction 0◦ and transverse direction 90◦. the indentation modulus in the longitudinal direction (0◦) is given by m0◦ = 2 √ c231 −c213 c11 ( 1 c44 + 2 c31 + c13 )−1 (3) where, c31 = √ c11c33. similarly, the indentation modulus in the transverse direction (90◦) is given as m90◦ = √ m12m13 (4) where, m12 = √ c11 c33 m3 and m13 = c211 −c212 c11 . the other well defined solution schemes for general anisotropic elastic materials are proposed by vlassak et al. [3]. the solution relates the indentation modulus with the elastic properties of the material as well as indentation direction. the indentation modulus for the anisotropic material according to vlassak et al. are given by m = 1 α (e, φ) (1 −e2) 14 (5) with the assumption of elliptical contact area formed with the indenter, the function α (e, φ) relates to eccentricity (e) of the contact ellipse and angle between the major axis of the ellipse and the reference direction of a function h (refer to 3 for further details). the solutions of vlassak et al. are used to determine the elastic constants by comparing the indentation modulus with the experimental results performed at varying orientation angles. the values of el, et, and glt are varied such that the solution matches the experimental indentation modulus closely. the values of νlt and νtt are kept constant as their effect on the indentation modulus is considered to be negligible [2]. similarly, the expressions of indentation modulus given by delafargue and ulm [7] are compared with the experiments to determine the elastic constants at two independent directions. 4. results and discussions from microscopic analysis of the carbon fibers i.e., according to section 2.2, we found that the actual angle of orientation is different than the cutting angle. for carbon fibers with 0◦ i.e., carbon fibers orientated in a longitudinal direction, there exists small deviation in the orientation angle, and the actual angle was found to be 12◦. but, for 90◦ practical orientation remains same as the cutting angle. also, it is found that there exists a small deviation in the orientation angle for the remaining cross-sectioned samples. the results of nanoindentation experiments of carbon fibers with varying orientation angle with its axis are presented. the load (p) versus the indentation 109 t.s. guruprasad, v. keryvin, a. bourmaud acta polytechnica 0 30 60 90 120 h [nm] 0.0 0.5 1.0 1.5 p [m n] 90 12 test_1 test_2 test_3 test_4 test_5 test_6 figure 3. load (p) versus indentation depth (h) behaviors from nanoidentation experiments for carbon fibers orientated at 12◦ (solid lines) and 90◦ (dotted lines). depth (h) from the indentation experiments are plotted for all the samples between 12◦ and 90◦. as the diameters of the carbon fibers indented are 5.2 µm and it is difficult to locate and perform indentation at the middle of the fibers. figure 3 shows the representative p −h curves of carbon fibers orientated at 12◦ and 90◦ with its axis. from the plot, we can see that there exists a repeatability in the p −h curves for two of those orientations and similar repeatability in the p − h are observed for other orientations as well. for the carbon fibers with orientation of 12◦, we observe from the p −h curves that there exists a hysteretic behavior and a small plastic deformation in the material but for the samples orientated at 90◦ shows negligible plasticity and are nearly elastic. the p −h also shows that the load required for the nearly longitudinal orientated fibers (12◦) requires approximately three times the load in comparison to carbon fibers oriented in transverse direction (90◦) to produce a same indentation depth. the indentation modulus (m) versus the actual orientation angle (θ) from the experiments are plotted in figure 4. due to anisotropy in the carbon fibers, the value of the indentation modulus decreases with increase in the orientation angle. apart from the indentation modulus, we also observed a similar decreasing behavior in the hardness with increase in the orientation angle. from the theoretical models presented in section 3, elastic constants are determined from the nanoindentation experiments. using vlassak et al. solutions, the elastic constants such as el, et, and glt are varied independently, such that a close match is obtained with the plot of experimental indentation modulus versus orientation angle. the final predicted values of three elastic constants are shown in table 1 and the poission’s ratio’s used are νlt = 0.3, νtt = 0.3. the corresponding fit with the experiments is shown in 10 10 30 50 70 90 [°] 0 20 40 60 80 100 m [g pa ] delafargue and ulm experiments vlassak et al. figure 4. the plot of indentation modulus (m) versus the orientation angle (θ) from nanoindentation experiments (circular markers), theoretical solutions of delafargue and ulm [7] (square markers) and vlassak et al. [3] (solid line). figure 4 and the obtained fit has a good agreement with the nanoindentation experiments. our predicted values of transverse modulus (et) and the shear modulus (glt) are comparable with other carbon fibers (type t-300), which was reported by csanádi et al. [4] and the values are et = 30.9 gpa, glt = 10.9 gpa. el (gpa) et (gpa) glt (gpa) 296 25.5 9 table 1. the predicted values of the elastic constants (el, et, and glt) of carbon fibers (im2c) via combined experimental-theoretical approach. similarly, the indentation modulus using the closedform solutions of delafargue and ulm are calculated for two orientations i.e., for 0◦ and 90◦ by varying the elastic constants. here, the results shows a good agreement with vlassak et al. as well as the indentation experiments. 5. conclusions the anisotropic elastic properties of carbon fibers (im2c) are determined using combined experimentaltheoretical approach from nanoindentation. the carbon fibers are cut at varying orientation angles between 12◦ and 90◦, and they are indented to determine the indentation modulus between longitudinal orientation and the transverse orientations. then, the plot of experimental indentation modulus versus the orientation angle is closely matched with the solutions of delafargue and ulm and also with the solutions of vlassak et al. to predict the elastic constants. 110 vol. 27/2020 anisotropy list of symbols el elastic modulus in logitudinal direction [gpa] et elastic modulus in transverse direction [gpa] glt shear modulus [gpa] νlt,νtt poisson’s ratios m indentation modulus [gpa] θ orientation angle [◦] cij elastic constants of the material p load [mn] h indentation depth [nm] references [1] r. gibson. principles of composite material mechanics. fourth edition. boca raton: crc press, 2016. doi:https://doi.org/10.1201/b19626. [2] a. jäger, t. bader, k. hofstetter, j. eberhardsteiner. the relation between indentation modulus, microfibril angle, and elastic properties of wood cell walls. composites part a: applied science and manufacturing 42(6):677 – 685, 2011. doi:https://doi.org/10.1016/j.compositesa.2011.02.007. [3] j. vlassak, m. ciavarella, j. barber, x. wang. the indentation modulus of elastically anisotropic materials for indenters of arbitrary shape. journal of the mechanics and physics of solids 51(9):1701 – 1721, 2003. doi:https://doi.org/10.1016/s0022-5096(03)00066-8. [4] t. csanádi, d. németh, c. zhang, j. dusza. nanoindentation derived elastic constants of carbon fibres and their nanostructural based predictions. carbon 119:314 – 325, 2017. doi:https://doi.org/10.1016/j.carbon.2017.04.048. [5] s. duan, f. liu, t. pettersson, et al. determination of transverse and shear moduli of single carbon fibres. carbon (158):772 – 782, 2020. doi:https://doi.org/10.1016/j.carbon.2019.11.054. [6] s. r. swanson. hertzian contact of orthotropic materials. international journal of solids and structures 41(7):1945 – 1959, 2004. doi:https://doi.org/10.1016/j.ijsolstr.2003.11.003. [7] a. delafargue, f.-j. ulm. explicit approximations of the indentation modulus of elastically orthotropic solids for conical indenters. international journal of solids and structures 41(26):7351 – 7360, 2004. doi:https://doi.org/10.1016/j.ijsolstr.2004.06.019. [8] w. oliver, g. pharr. an improved technique for determining hardness and elastic modulus using load and displacement sensing indentation experiments. journal of materials research 7(6):1564–1583, 1992. doi:10.1557/jmr.1992.1564. [9] w. oliver, g. pharr. measurement of hardness and elastic modulus by instrumented indentation: advances in understanding and refinements to methodology. journal of materials research 19(1):3–20, 2004. doi:10.1557/jmr.2004.19.1.3. 111 https://doi.org/https://doi.org/10.1201/b19626 https://doi.org/https://doi.org/10.1016/j.compositesa.2011.02.007 https://doi.org/https://doi.org/10.1016/s0022-5096(03)00066-8 https://doi.org/https://doi.org/10.1016/j.carbon.2017.04.048 https://doi.org/https://doi.org/10.1016/j.carbon.2019.11.054 https://doi.org/https://doi.org/10.1016/j.ijsolstr.2003.11.003 https://doi.org/https://doi.org/10.1016/j.ijsolstr.2004.06.019 https://doi.org/10.1557/jmr.1992.1564 https://doi.org/10.1557/jmr.2004.19.1.3 acta polytechnica 27(0):107–111, 2020 1 introduction 2 experiments 2.1 materials 2.2 microscopic analysis 2.3 nanoindentation 3 theoretical models for anisotropic nanoindentation 4 results and discussions 5 conclusions list of symbols references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2021.31.0010 acta polytechnica ctu proceedings 31:10–17, 2021 © 2021 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague research tools aplicable in designing of high-speed and high-power rail vehicles vojtěch dybala czech technical university in prague, faculty of mechanical engineering, department of automotive, combustion engines and railway engineering, technická 4, prague, czech republic correspondence: vojtech.dybala@fs.cvut.cz abstract. the importance of the railway transport of both goods and passengers continuously grows as it is in many points of view more ecological and economical solution in comparison with road transport. just the importance has been supporting efforts to design more powerful and faster trains reaching traction powers more than 1.6 mw per a wheel-set or service top speed much more than 300 km/h till these days. to reach design which has enabled such a kind of performances it was necessary to research dynamic behaviour of railway vehicles. both via laboratory measurements and simulations. the laboratory experiments have been carried out on a specially designed laboratory equipment called roller rigs. a laboratory equipped by roller rig for testing of the railway vehicle dynamic behaviour has been built at the faculty of mechanical engineering at ctu in prague, also within phd study programs and sgs grants. another powerful tool within research activities is a simulation. kind of a such simulation will be presented by this contribution with the focus on the torsion dynamics of high-power fully-suspended drive of a railway vehicle, which has been developed also by phd students under financial support of the faculty of mechanical engineering and related grants. keywords: railway vehicle, dynamic, roller rig, rollers, drive, torsion dynamics. 1. introduction in the railway vehicle’s research has been used different tools. these are calculations and simulations carried out via specialized software, measurements on real vehicles on real tracks and measurements carried out in laboratories on roller rigs. the importance of roller rigs comes just from laboratory conditions and its advantages. the possibility to define exact conditions of measurements, ability to reproduce measurements, research of dangerous phenomena and low costs against measurements on real vehicles on real tracks. all of this caused from the 2nd half of the 20th century till these days an expansion of roller rigs in connection with the need to research dynamic behaviour of railway vehicles. whereas roller rigs were built for base research there has been also a need to do an applied research on the engineering level via simulations in the industry. for these purposes various software has been developed and used. the university setting enables to apply both approaches for phd students, however its limited in comparison with industry companies and specialized research centres. nevertheless, applicability of roller rigs and simulation results of torsion oscillations and their excitation were achieved. 2. roller rigs the roller rig is a device, which consists of three main parts. these are a frame or an external construction of the roller rig, rollers and a propulsion system of rollers with related components as e.g. couplings and shafts. the most important are the rollers as these components shall simulate rails as a part of the railway infrastructure. the rollers simulating endless rails can be of two basic design types, as applied for roller rigs at ctu in prague: • disc type rollers – figure 1 left, • drum type rollers – figure 1 right, which is a prototype roller from the historical point of view roller rigs touches the remote past. for example [1], figure 2 presents the roller rig from swindon in england specified to measure a traction power of steam locomotives. another example of a roller rig, which was built in the real-life scale, can be the workplace created in 1977 by deutsche bahn ag in munich, see [1, 2] and figure 3. this research workplace was dedicated for the research and development of the first generation of ice high-speed train. this device enabled to research train dynamic behaviour on its limits under different conditions – speeding in the straight track, acceleration and braking behaviour or ride on a track with deviated geometry etc. except of these big real-life scale devices reduced scale devices have been created. in this area the scaled roller rigs, which has been built at the faculty of mechanical engineering at ctu in prague, see figure 4 and figure 5, can be taken as examples. the first roller rig has been used for the education and its abilities and continuous development were presented [3, 4]. student’s measurements are focused on the transverse dynamic of railway bogie movements within the education. 10 https://doi.org/10.14311/app.2021.31.0010 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 31/2021 designing of high-speed and high-power rail vehicles figure 1. roller rigs – faculty of mechanical engineering at ctu in prague. 1 figure 2. the roller rig in swindon [1]. the second roller rig has been under construction and without a practical used till these days. its purpose is generally oriented into torsion dynamics and for first attempts to measure the adhesion phenomena in the wheel-roller contact as a simulation of the wheel-rail contact, which we know from numbers of measurements. as a good example can be used results presented in [5] and figure 6. the basic idea for the supposed measurements of the adhesion coefficient comes from the picture of forces, which act in the wheel-rail contact as presented in [7] and figure 7 and the knowledge, that the adhesion coefficient is a function of a wheel slip. according to the equation (1) the tangential force t1, which is an equivalent of the wheel torque and vertical wheel force q will be measured and adhesion coefficient µ will be calculated. t1 = qµ (1) as the adhesion coefficient is a function of the wheel slip, the slip itself will be calculated based on the difference in rotation speed of a wheel-set and a roller, which will be measured. on the real vehicle the slip is defined based on the difference of the vehicle speed and real rotation speed of a wheel-set. this fact is defined by the equation (2) and its deduction in [7] and figure 8. s = rkωk − v v (2) 3. torsion oscillations research within a generally defined task to research transition and electromechanical phenomena in traction drives of railway vehicles, the phenomenon of torsion oscillations has been researched as well. the torsion oscillation problems were studied also in the past, which is documented in [5]. that contribution presents results from simulations of the torsion dynamics for the mathematical model of the experimental roller rig with different types of wheelset drives. this contribution deals with torsion dynamics and oscillations, which are excited by harmonic components of the traction motor electromagnetic torque. the intended simulation model, which was created in matlab simulink, utilizes the idea, that the railway vehicle can be simplified into three basic separate sub-models from the perspective of the drive system. these are the electrical part, the mechanical part and the wheel-rail contact part, as shown in [8] and figure 9. fundamentals and function descriptions of the model as a whole were published in [7–9] and therefore no attention will be payed to them. the intention of this contribution is to present results of simulations focused on the oscillations. the phenomenon of torsion oscillations has been researched via a mathematical model which is equivalent to a fully-sprung mechanical drive of a locomotive, see [7] and figure 10. for the simulation model the drive has been reduced into a schematic torsion system as presented in figure 11. particular rotation masses, defined by their mass of inertia jx and torsion springs between them, defined 11 vojtěch dybala acta polytechnica ctu proceedings figure 3. the roller rig built by db-ag in munich [1, 2]. figure 4. the roller rig – faculty of mechanical engineering at ctu in prague. by their torsion stiffness kx are mathematically represented within the simulation model like in [7] and figure 12. the electromagnetic torque m produced by the asynchronous motor model, which accelerates whole model of the railway vehicle, is the source of excitation of the oscillations in the torsion system, because it consists of a nominal value and superimposed harmonic components. the course of this rippled torque within a simulation can be seen in figure 13. fast fourier transformation (fft) of the electromagnetic torque in figure 14 revealed harmonic components with typical frequencies. these harmonic components interact with the mechanical system in such a way, that they excite the oscillations. simulations proved, that excited oscillations can reach significant magnitudes for oscillation frequencies, which are the natural frequencies of the mechanical system. as in this simulation case dominant oscillations for frequencies around 2400 hz appeared in figure 5. the roller rig – faculty of mechanical engineering at ctu in prague. figure 6. comparison of measured and calculated coefficient of adhesion [6]. fft in gear’s acceleration torques, the example for the gearwheel acceleration torque course is shown in figure 15 and figure 16. table 2 shows that the magnitude of harmonic 12 vol. 31/2021 designing of high-speed and high-power rail vehicles harmonic frequency [hz] harmonic component magnitude in relation to nominal value of signal [%] phenomenon description 137 106 3rd multiple of 1st harmonic of motor supply voltage 800 and its multiples 10 to 55 even and odd multiples of inverter switching frequency +/137 0 to 8 side bands of switching frequency multiples table 1. harmonic frequencies of electromagnetic torque – overview and description. harmonic frequency [hz] magnitude of harmonic component in relation to nominal value of signal [%] phenomenon description 2400 105 3rd multiple of inverter switching frequency 2401 185 7th natural frequency of the mechanical system table 2. harmonic frequencies of the gearwheel accelerating torque – overview and description. figure 7. wheel-rail contact forces [7]. components of the rippled torque can be significantly higher than nominal value of the torque. in presented case 85% higher for the frequency of 2401 hz. this behavior is caused by the fact, that the 3rd multiple of the inverter switching frequency 2400 hz contained in the electromagnetic torque, which excites system’s oscillation, is close to 7th natural frequency of the figure 8. wheel slip deduction [7]. figure 9. visualization of fundamental model parts [8]. mechanical system ≈ 2401 hz. 4. conclusions with respect to the roller rigs as research and education tools at the faculty of mechanical engineering, ctu in prague the intention is to build the roller rig, which will enable to realize measurements of adhesion 13 vojtěch dybala acta polytechnica ctu proceedings figure 10. locomotive fully-sprung drive – cross section [7]. figure 11. schematic torsion system. 14 vol. 31/2021 designing of high-speed and high-power rail vehicles figure 12. mass and spring mathematical representation [7]. coefficient in the first step and afterwards possibly extend laboratory experiments for students. regarding nowadays results from the research oriented towards the torsion oscillations in the railway vehicle traction drives we can say, that these oscillations with significantly high magnitudes can be observed within torques loading particular components of the traction drive even in steady states, not only during transition phenomena. however, such observed torque ripple does not lead to the overloading of these components with respect to their strength, it will have an impact on the damage cumulation and thus on their life parameters. this idea is heading this research to be focused on the investigation of sensitivity of different traction drive’s design arrangement on the torsion oscillations and possibilities how to reduce them by means of utilization of modified standard components as couplings or gears. 5. list of symbols fx coefficient of friction [-], t1 tangential force [n], q vertical wheel force [n], µ coefficient of adhesion [-], s wheel slip [%], rk wheel radius [m], ωk angular speed of a wheel [rad/s], ν vehicle velocity [m/s], jx mass of inertia [kg·m2], kx torsion stiffness [n·m·rad−1], ϕx angle [rad], r1 pinion rolling radius [m], r2 gearwheel rolling radius [m], i gear ratio [-], m motor torque, elmag. [n·m], ω1,ω2 angular speed [rad·s−1], m1,m2 torque [n·m]. acknowledgements this research has been realized using the support of the technology agency of the czech republic, programme national competence centres, project #tn01000026 josef bozek national center of competence for surface transport vehicles and the support of the czech technical university and the related grant – no. sgs20/120/ohk2/2t/12. this support is gratefully acknowledged. references [1] n. bosso, a. gugliotta, a. soma. simulation of narrow gauge railway vehicles and experimental validation by mean of scale tests on roller rig, 2008. online, issn 15729648, https://doi.org/10.1007/s11012-008-9128-4. [2] a. jaschinski, h. chollet, s. iwnicki. the application of roller rigs to railway vehicle dynamics. vehicle system dynamics 31, amsterdam, 1999, https://doi.org/10.1076/vesd.31.5.345.8360. [3] s. myamlin, j. kalivoda, l. neduzha. testing of railway vehicles using roller rigs. procedia engineering 2017, 2017, issn 1877-7058, https://doi.org/10.1016/j.proeng.2017.04.439. [4] j. kalivoda, p. bauer. mechatronic bogie for roller rig tests. the dynamics of vehicles on roads and tracks – proceedings of the 24th symposium of the international association for vehicle system dynamics, 2016, https://doi.org/10.1201/b21185-96. [5] j. kolář, m. dub. matematické modelování torzní dynamiky pohonu dvojkolí. in současné problémy v kolejových vozidlech 2017: xxiii. konference s mezinárodní účastí, sborník příspěvků, pardubice, 2017, p. 195-204, isbn 978-80-7560-085-1. [6] o. polach. creep forces in simulations of traction vehicles running on adhesion limit. 6th international conference on contact mechanics and wear of rail/wheel systems, göthenburg, june 10-13, 2003. [7] v. dybala, m. libenský, b. šulc, c. oswald. slip and adhesion in a railway wheelset simulink model proposed for detection driving conditions via neural networks. sborník vědeckých prací vysoké školy báňské – technické univerzity ostrava, řada strojní, ostrava, 2018, issn 1804-0993 (online), https://doi.org/10.22223/tr.2018-1. [8] v. dybala. simulace zvlnění elektromagnetického momentu asynchronního trakčního motoru lokomotivy napájeného ze střídače. konference studentské tvůrčí činnosti 2019, sborník konference, praha, 2019, isbn 978-80-01-06564-8. [9] v. dybala. vliv regulace na frekvenční složení fázových proudů a elektromagnetického momentu při simulování napájení asynchronního motoru ze střídače. súčasné problémy v koľajových vozidlách – prorail, 2019. diel i, žilina, 2019, isbn 978-80-89276-58-5. 15 https://doi.org/{10.1007/s11012-008-9128-4} https://doi.org/{10.1076/vesd.31.5.345.8360} https://doi.org/{10.1016/j.proeng.2017.04.439} https://doi.org/{10.1201/b21185-96} https://doi.org/{10.22223/tr.2018-1} vojtěch dybala acta polytechnica ctu proceedings figure 13. rotor speed and rippled torque of asynchronous motor. figure 14. fft analysis of the electromagnetic torque. 16 vol. 31/2021 designing of high-speed and high-power rail vehicles figure 15. course of gearwheel accelerating torque. figure 16. fft analysis of the gearwheel accelerating torque. 17 acta polytechnica ctu proceedings 31:10–17, 2021 1 introduction 2 roller rigs 3 torsion oscillations research 4 conclusions 5 list of symbols acknowledgements references https://doi.org/10.14311/app.2022.33.0504 acta polytechnica ctu proceedings 33:504–510, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague how design concepts influence carbon footprints of load bearing structures anne rønninga, ∗, kjersti prestruda, simon saxegårda, simen s. haaveb, magne lysbergb a norsus norwegian institute for sustainability research, stadion 4, 1671 kråkerøy, norway b arcon prosjekt as, verftsgata 11 pb. 224, 7801 namsos, norway ∗ corresponding author: anne@norsus.no abstract. in recent years, it has become essential to consider the total carbon footprint of a construction project. commonly, the question has been: ’what is the best material to be used in this context?’ in this paper we argue that this question is incomplete, not taking the complexity of design choices into consideration. this paper intends to share light on how to analyse some factors that influence the construction of buildings in order to contribute to climate change mitigation, taking this complexity into consideration. calculation of fossil greenhouse gas (ghg) emissions for two load-bearing structures for office buildings in 4, 8 and 16 storeys with equal functional requirements; e.g. load bearing capacity, acoustic performance, fire resistance and adaptability are addressed. the main materials for the load-bearing structures are cross laminated timber (clt) elements and precast concrete elements respectively. the result show that one cannot on a general basis conclude that either type of load-bearing structure cause less fossil ghg emissions. it is always important to consider the building design, functionality as well as external conditions such as location when considering different load-bearing structure materials. keywords: carbon footprint, epd, lca. 1. introduction over the last 30 years, life cycle assessments (lca) have been applied in the construction sector as the methodological foundation to evaluate environmental performance of construction works and materials over the life cycle [1–3]. standards for both product and building level assessment are developed. for the latter, the european standard en 15978:2011, specifies the calculation method for assessing the environmental performance of building [4]. for documentation of the environmental profile for construction products, en 15804 provides the requirements to develop environmental product declarations (epd) [5]. the system boundaries defined in en 15978 are a1-a3 (product stage), a4-a5 (construction process stage), b1-b7 (use stage), c1-c4 (end of life stage) and d (benefits and loads beyond the system boundary). the approach covers all stages of the building’s life cycle and is based on data obtained from epd, their information modules, and when appropriate other information necessary and relevant for carrying out the assessment of the environmental performance of the building. to meet the increasing demands for ghg calculations of buildings in the norwegian market, a standard, ns 3720:2018 that define the rules and requirements for calculation of so, is published by standards norway [6]. moncaster et al. reviewed and analysed the data from over 80 individual life cycle assessments of buildings [7]. they found that several authors have used such findings to identify routes to lower carbon buildings. the strategy considered to have the biggest impact is the substitution of high carbon materials with low carbon; this is considered particularly important for the main structural and cladding elements which are often shown to have the highest impacts [8–10]. others have found that adaptability and patterns of use are of more importance as design factors than the building materials and products themselves, as the latter are a consequence of those factors [11, 12]. additionally, high replacement rates of materials with high embodied carbon as consequence of low adaptability will have a great impact on life cycle performance. the average rental period for office buildings in norway is 7 years. every seventh year, buildings are extensively rebuilt due to new tenants’needs and requirements [1]. how extensive the rebuilding processes will be will depend on the degree of adaptability of the building. thus, design for adaptability for office buildings is of vital importance. the goal of this study is to increase the knowledge regarding how design for adaptability, choice of materials and locations affect greenhouse gas calculations for load-bearing structures in office buildings. 504 https://doi.org/10.14311/app.2022.33.0504 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 carbon footprints of load bearing structures figure 1. illustration of the two different load-bearing structures for office buildings (wood-based structure to the left). 2. methods 2.1. the objects the objects under study are two load-bearing structures for office buildings in 4, 8 and 16 storeys respectively, with equal functional requirements; e.g. load bearing capacity, acoustic performance, fire resistance and adaptability for future changes.the design is similar for both the wood-based and the concrete based structure. a visualisation of the two alternative 4 storey load-bearing structures can be seen in figure 1. the floor area size per storey of 50.4 m × 15.9 m (801 m2) was chosen based upon recommendations given in sintef building research design guides for adaptable office buildings [13]. it includes guidelines for physical sizes such as building width and floor area as well as guidelines for flexibility in office buildings in connection with pillar locations, etc. a width of 15−17 m is optimal as it gives the possibility of different interior solutions and simultaneously an acceptable area efficiency [14]. ideally, columns and load-bearing walls should be avoided in areas that can accommodate different types of workplace solutions to ensure high degree of adaptability. based on building technical considerations, the optimal span for a building width of 16 m will be approx. 8 m. this means that a row of columns should be placed approximately in the middle of the building with a centre distance in the longitudinal direction of approx. 5 − 8 m. this recommendation is to limit the load and cross-sectional dimension of the beams, between the columns and the column itself. it is assumed that foundations for all alternatives consist of the same concrete quality b35, placed in a humid environment below ground level. finally, the soil type is assumed to be natural sandy for 4 storeys, and natural gravel for 8 and 16 storeys. the solution of the foundation in the study consists of a foundation slab at the two short sides. the constructions experience the same magnitude of horizontal forces from wind, but the dynamic effect will be different since the stiffness and self-weight of the two alternative 16 storey buildings is different. the gables in the concrete solution are assumed to be of 12 m width, consisting of vertical, rigid concrete slabs, whereas the two short sides in the woodbase structure are assumed to consist of trusses along the width of the building (15.9 m). volume of concrete walls in every gable is 126 m3 (stabilizing effect from self-weight is about 2800 kn), whereas volume of trusses is 46 m3 (stabilizing effect from self-weight is about 190 kn). this explains the increased need for ready-mixed concrete for 16 storey wood-based structure. the fire requirements in norwegian building regulation, tek17 is used [15]. the regulations consider the constructions to be in hazard class 2, independent of the number of storeys, given that all storeys are used as office space. fire design of the concrete solution would be unproblematic up until r120. this does not apply to wood-based constructions which have to meet requirements on design and measures. proposed measures are (≥ r 60) 12 mm gypsum and 2 layers of 14 mm fire rated gypsum [16]. the load-bearing structures are designed by use of fem-design by strusoft and os-prog and quantities are extracted from this and further organised in excel, see table 1 and table 2 for material specification and quantities. 2.2. system boundaries en 15978 [4] and ns 3720:2018 [5] was used as the underlying rules for carrying out the calculation of the ghg emissions for the load-bearing structures. several of the applied epds that are used as data source have not declared emissions connected to the c module (end-of-life). therefore, it was chosen to limit the study to include only a1-a5 modules, from the extraction of raw materials to the construction of the load bearing structures. building shell and façade system, non-bearing separating walls, or technical equipment such as ventilation, heating, sanitary 505 a. rønning, k. prestrudr, s. saxegård et al. acta polytechnica ctu proceedings construction part material 4 storeys 8 storeys 16 storeys unit epd number foundation ready-mixed concrete 423 556 2 713 tonne nepd-332-216-no reinforcement 17 22 106 tonne nepd-347-238-en columns − − 31 m3 nepd-1577-605 beams − − 25 m3 nepd-1577-605 clt-slabs − − 203 tonne nepd-345-236-no vertical structures columns 23 72 298 m3 nepd-1577-605 truss 11 23 109 m3 nepd-1577-605 fireproof plasterboard − 20 000 41 550 m2 nepd-1264-406-en plasterboard − 10 000 20 775 m2 nepd-1260-406-en steel plate 2.95 8.45 37.9 tonne nepd-402-281-en bolts (steel) 3.29 9.22 41.9 tonne nepd-402-281-en horizontal structures beams 99 198 397 m3 nepd-1577-605 clt-slabs 813 1 626 3 252 m3 nepd-345-236-no steel plate 3 6 11 tonne nepd-402-281-en floor plasterboard 3 100 6 250 13 500 m2 nepd-110-177-en insulation 222 517 1 182 m3 nepd-1696-683 elastic underlayment 3 100 6 250 13 500 m2 nepd-207-260-no wooden laths 18 43 99 m3 nepd-308-179 table 1. material amounts for the clt-based load-bearing structure. construction part material 4 storeys 8 storeys 16 storeys unit epd number foundation ready-mixed concrete 476 670 1 583 tonne nepd-332-216-no reinforcement 19 26 78 tonne nepd-347-238-en columns − − 79 tonne epd generator beams − − 62 tonne epd generator hollow core slabs − − 204 tonne epd generator vertical structures columns 47 124 565 tonne epd generator steel plate 0.646 1.26 2.54 tonne nepd-402-281-en bolts (steel) 2.07 4.15 8.29 tonne nepd-402-281-en wall panels 102 409 1292 tonne epd generator reinforcement reinforcement 13 42 98 tonne nepd-326-206-enprestressed reinforcement 14 29 57 tonne nepd-458-296-en horizontal structures beams 250 499 999 tonne epd generator insulation 3100 1634 13 500 m2 nepd-00131e_rev1_rockwool hollow core slabs 817 6250 3267 tonne epd generator table 2. material amounts for the pre-cast concrete-based load-bearing structure. 506 vol. 33/2022 carbon footprints of load bearing structures figure 2. fossil ghg emissions for three alternative load-bearing structures located in kristiansand and trondheim given as kg co2e/m2 and electricity are not included, as it were considered independent from the load-bearing systems. the alternative structures are dimensioned for the same service life, being maintenance-free during the service life and that they will not affect the operational conditions of the buildings. thus, the use phase (b1-b7) is excluded. no assessment has been made of how large the contribution from the module c is, but erlandsson and malmqvist [17] demonstrated in their study that fossil greenhouse gas emissions related to demolition and waste management of concrete and wood-based buildings were in the order of 2 − 5 % of emissions from a1-a5. carbon dioxide that is absorbed over time as the tree grows, is stored in wood-based building products during the life of the building [18]. when the buildings are disposed of after the end of their life and the wood-based products are used for energy recovery, the carbon is emitted in the form of co2. the climate change of sequestering biogenic carbon, storing it in harvested wood products and substituting more emission-intensive materials are hard to quantify. although different methodological choices and assumptions can lead to different conclusions, there is no consensus on the assessment of biogenic carbon in life cycle assessment [19]. as we have not included end-of-life assessment, instant oxidation of biogenic carbon is used as the approach where biogenic carbon is considered "climate neutral". this is in accordance with the common practice with calculations that do not include all life cycle modules, and with the practice for ghg calculations for constructions [4, 18]. 2.3. environmental data epd for construction products published by epdnorway are used as data source for the ghg emissions for the building materials specified in table 1 and table 2. the amount of all materials used for the four-, eightand sixteen-storey load-bearing structures are given in table 1 and table 2 for cltand pre-cast concrete structure respectively. epd for construction products published by epd-norway are used as data source for the ghg emissions for the building materials specified in table 1 and table 2. where more than one epd for the same representative product was available, the product with lowest ghg emissions are used. emissions related to the a4 are usually calculated based on an average distance from the production site to a typical construction site in norway, or to a specific destination such as oslo. this means that the assumptions for a4 will vary and the values in the epd cannot be applied directly. in this study, the construction sites are kristiansand and trondheim. ghg emissions were calculated based on the same type of means of transport as stated in the epd. our transport calculations also include infrastructure related to transport, which is not necessarily the case for all a4 modules in the epd. the precast concrete manufacturers in norway use a pre-verified epd-generator that allows for producing project specific epds [20]. two different data set are used; average data from four different precast manufacturers’epd for columns, slabs, wall elements and hollow cores respectively named ’typical element recipes’and epd for the same product from the manufacturer epd with the lowest ghg emission for a1a3. for clt slabs the manufacturer with lowest documented ghg emissions for a1-a3 is chosen. 3. results in figure 2 ghg emissions per square metre for the wood-based and concrete load-bearing structures are given for all three storey heights located in kristiansand and trondheim respectively. the results show that the wood-based structure has lower fossil ghg emissions than the construction with pre-cast concrete elements when built on four storeys for both locations but can be reversed at 16 floors when using the best precast concrete elements. for all cases, materials used for both vertical 507 a. rønning, k. prestrudr, s. saxegård et al. acta polytechnica ctu proceedings kristiansand trondheim element 4 storeys 8 storeys 16 storeys 4 storeys 8 storeys 16 storeys clt-slabs 43 % 34 % 25 % 36 % 28 % 21 % ready-mixed concrete incl. reinforcement (foundation) 28 % 15 % 26 % 31 % 16 % 27 % fire measures (plasterboard) 6 % 25 % 19 % 8 % 27 % 20 % steel products 11 % 12 % 16 % 13 % 13 % 17 % table 3. share of fossil ghg emissions in percentage from different construction materials in wood-based loadbearing structures of 4-, 8and 16-storey located in kristiansand and trondheim. kristiansand trondheim element 4 storeys 8 storeys 16 storeys 4 storeys 8 storeys 16 storeys hollow core slab incl. reinforcement 37 % 37 % 32 % 37 % 37 % 31 % ready-mixed concrete incl. reinforcement (foundation) 28 % 20 % 20 % 28 % 19 % 45 % columns/beams 20 % 18 % 20 % 18 % 16 % 46 % reinforcement 7 % 9 % 9 % 10 % 12 % 12 % wall element 6 % 12 % 15 % 5 % 10 % 15 % table 4. share of fossil ghg emissions in percentage from different construction materials in concrete load-bearing structures of 4-, 8and 16-storey located in kristiansand and trondheim. and horizontal structures contribute most. table 3 and table 4 present the contribution analysis to the total fossil ghg emissions in more details. the contribution analysis shows that clt-slabs, and foundation (ready-mixed concrete incl. reinforcement) contribute most to ghg emissions for the wood-based structure. the share of emissions from clt decreases with the height of the building. due to need for fire protection at higher storey, the contribution from plasterboard is significant. for the 8and 16-storey structure, fire measures contribute with 25 % and 19 %, and 27 % and 20 %, of the total emissions for structures located in kristiansand and trondheim, respectively. furthermore, ghg emissions from transport of clt elements from the manufacturer in sweden to kristiansand accounts for 51% of the total ghg emissions from clt (49 % for a1a3). the distance to trondheim is shorter, and the transport contribution from clt is then 36%. this illustrates that transport to building site for heavy materials can be important for the results and is conformed by [21]. for the pre-cast concrete structure, hollow core slabs and columns/beams are the elements that contribute with the highest share of ghg emissions. also, for the heavy hollow core slabs, transport turned out to be significant. the manufacturer that produce the hollow core slabs with significant lower ghg emissions than the competitors, is chosen only for deliveries to kristiansand. the transport from production site to trondheim gives ghg emissions that makes manufacturers closer to trondheim favourable. 4. discussion the span of the construction in this study is set to 7.95 m. for solid wood constructions a span of 5 to 5.5 meters is an alternative to avoid dimensions that are too large. if the span length is reduced, the thickness of the slabs is reduced including column and beams dimensions, but there will be a need to add an extra load-bearing axis. thus, the total material amount may not be significantly changed. nevertheless, the structure will have a lower adaptability for changes and consequently lead to more materials e.g. when new tenants every seventh year require changes in the interior [2]. the gross area is representative for office buildings in norway [13]. when assessing the higher structures, the choice of shape implies constraints for how the foundation and fire protection is solved. the foundation design is affecting the choices of materials and the material amounts as the structures. as the higher structures are relatively tall and slender (16 storey) the foundation will be subjected to large tensile forces. this can be solved by two different approaches; by friction piles or as in this study, by a casted underground concrete plate to ensure enough deadload to reach equilibrium. as the tensile forces are greater for the wood-based structures, the amount of materials in foundation used to ensure equilibrium are greater, which explains the ghg emissions (26 and 26 %) for the 16-storey wood-based structures. it is worth mentioning that there are also different solutions as to how the foundations of this specific construction design can be structured. a construction that has a different shape, on the other hand, 508 vol. 33/2022 carbon footprints of load bearing structures will have different characteristics which will be best solved by utilizing a different approach due to the shape (tall and slender) of the higher structures, the building will have a propensity to oscillate due to wind forces. this is not taken into consideration in our study but may be a challenge for the highest wood-based structure. a possible remedy to satisfy the requirements on oscillation is to build weight and masses into the floors for the woodbased structure. however, this again will require more columns or increased dimensions for columns and beams that will affect the amount of material used. today’s available technology for fire safety in wood constructions is assumed. as the technology develops on this area, the future solutions might be better and less material consuming, and it can potentially lead to a lower ghg emissions from the fireproof plasterboard used in the wood construction or solved by other technologies. however, research in the field of fire safety of clt buildings concludes that there is a lack of knowledge for use of unprotected clt [22, 23]. the fossil ghg emissions are highly influenced by transport and the choice of supplier used, as the elements for both the woodand concrete-based structure are quite large. the shape of the elements leads to more transport than other materials. this shows a considerable potential for reduction of ghg emissions especially for large elements, and it is important to include transportation when requirements for ghg emissions are set. it is pointed out that the fossil ghg emissions from material production and transport alone is not sufficient to conclude what is the most climate and environmentally friendly. the results of this study show that how the load-bearing structures are designed will considerably affect the ghg emissions related to both the load-bearing structure itself and the ability for design for adaptability for future changes in the office building. we want to emphasize in the study that based on the context and different solutions that are applied in specific projects, it will be possible to reduce the ghg emissions regardless of the material choices. this can be further promoted by giving manufacturers and other actors more room to use their skills and practical knowledge to develop innovative solutions. giving the designers the opportunity to adjust the design of a construction in collaboration with manufacturers, will make it possible to develop innovative solutions and potentially a lower carbon footprint. the carbon footprint of a construction project will rely on the design and the materials chosen for the building, which in turn will affect the future need for materials for maintenance and lay premises for how extensive future renovation or refurbishment will be. can any general conclusions be drawn based on the results of this study? what is considered the best material or design choices with respect to climate change mitigation will be contextual for each unique construction. and even when assessing a load-bearing structure and not the whole building, a holistic approach is vital and one need to reflect on how the load-bearing structure will affect other activities’potential ghg emissions during the service life to ensure design for low ghg emissions. references [1] m. röck, m. r. m. saade, m. balouktsi, et al. embodied ghg emissions of buildings the hidden challenge for effective climate change mitigation. applied energy 258, 2020. https://doi.org/10.1016/j.apenergy.2019.114107. 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[23] n. k. reitan, k. l. friquin, r. f. mikalsen. challenges concerning cross laminated timber in tall buildings. rise fire research, trondheim, 2019. 510 http://www.bco.org.uk/research/publications/bcoguidetospec2019.aspxhttps://www.treteknisk.no/resources/filer/publikasjoner/teknisk-handbok/hefte-4-brann.pdf https://byggteknikforlaget.se/olika-byggsystem-av-betong-och-tra-dar-mix-av-material-inklusive-stal-ger-klimatfordelar/ https://doi.org/10.3390/su10062020 acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0126 acta polytechnica ctu proceedings 26:127–133, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app mechanical testing and numerical modelling of porous structures improving oseintegration of implants petr vakrčka∗, aleš jíra, petra hájková czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: petr.vakrcka@fsv.cvut.cz abstract. implants, such as dental, are ordinary devices in medical care nowadays, even though they are quite expensive. in the present study, the use of trabecular and gyroid structures as external layer of implants is examined. the advantage of porous structures compared to surface modification of compact implants is the possibility to be fabricated by additive manufacturing together with the whole implant. the additive manufacturing also allows us to produce various shapes with controlled porosity for bone ingrowth. the design of 6 types of trabecular and 4 types of gyroid structures is part of the study. the trabecular structures are strut-based, whereas the gyroid structures are based on a wall system. the study is focused on mechanical testing of samples which were 3d printed from the titanium alloy ti6al4v. the gyroid structures, which we evaluated as more reliable, were chosen for numerical modelling. other observed advantages and disadvantages of the structures are also discussed. keywords: implant, cancellous bone, trabecular structure, gyroid, titanium alloy, additive manufacturing, numerical model. 1. introduction thanks to modern medicine and advancing technological progress, people across the world live longer than ever before. the technology has become vital for giving patients the best care as possible. many patients in the past had to suffer because there was no technology which could support them. nowadays, people do not have to suffer anymore because technology is on their side. since engineers started improving medical care, new medical devices have been created, modern software has been adapted, and technology has increase hospitality for every patient. for example, laparoscopic surgeries have become the standard for many operations. they are minimally invasive, cause less pain, and shorten the recovery time. in recent years, the next great achievement in the field of medicine is targeted cancer therapy, which can interfere with the spread of cancer by blocking cells involved in tumor growth. worth mentioning is also czech plastic surgeon bohdan pomahač, who performed the first full face transplant, and the list of recent achievements in the medical field goes on. however, the presented study focuses on the particular field of medicine which is implantology. the evolution made the body as perfect as possible. that is why the attempt to adequate reconstruction has been close to pursuing the holy grail. for this reason, the majority of attempts have been unsuccessful. the human effort has made massive progress since the first primitive implants from ancient egypt until custom-made implants made from titanium alloys by 3d printers. at first, the implants meant to be for privileged ones. nowadays, implants have become a device which is available for the ordinary people. even now, there are many challenges for engineers. the crucial topic is the contact area between an implant and living tissue. the presented study sets a goal to knock down the problem by using a different approach than most of the current implants. in general, commonly used compact implants need various surface modifications for adequate oseintegration. thanks to additive manufacturing, we can print almost every shape using a 3d printer. it allows us to substitute the surface modification with a suitable structure which can minimize the risk of implant loss. trabecular and gyroid structures, which can be seen all around ourselves, can be the answer to this complicated question. figure 1 shows examples of presence of the structures in nature. 2. geometrical model the main aim of the study is to design an artificial model of 3d-printed microstructure with similar properties as cancellous bone. it could allow us to design an implant that does not need a surface modification. we substitute the surface modification with porous micro-structures. this method has a variety of reasons. first of all, there is no universal approach of surface modification suitable for all biomaterials, but the porous structures can be used for all additive manufactured products. the second huge advantage is the control of pore size. this gives us a chance to 126 https://doi.org/10.14311/app.2020.26.0126 https://ojs.cvut.cz/ojs/index.php/app vol. 26/2020 title figure 1. various application of trabecular structures in nature [1]. in the top we can see the structure of leaf. in the bottom, we can see trabecular bone morphology [2]. optimize the pore size to maximize the bone ingrowth. in the end, if we change the porosity of the microstructure, we can modify young’s modulus to minimize the stress shield effect which can lead to resorption of the human bone. at first, trabecular (strut-based) structures were selected. after designing and fabrication of the samples, the microstructure was observed due to problematic different cooling rate which is caused by placing one layer upon the other during fabrication process. thin bars seemed to be prone to fracture due to cooling process of the structures. these fractures can split in time and lead to necrosis or aseptic loos of implant. it was assumed that the change of morphology could give us more suitable structure and the problem could be eliminated. what is more, the strength of trabecular structures was not as high as it had been expected, probably due to the fractures. this was the main reason why structures based on a wall system were chosen for mechanical testing and numerical modeling in the next step. namely, the gyroid structures were selected for its constant pore size throughout the model. so the next goal of this study was to design 4 gyroid structures with different thickness of the walls and size of the pores and to test the structures mechanically for verification of the numerical models. gyroid is a continuous and triply periodic cubic morphology with a constant curvature surface across a range of volumetric fill fractions. it can be found in a variety of natural and synthetic systems [3]. the most convenient approximation of the gyroid structure for our purpose is through a level surface. it is defined by the function: f(x,y,z) = t , (1) where f determines the form of the surface via spacedividing function and t is the constant which determines the volume fractions of the divided space. if we set the parameter t to zero and substitute the f with an appropriate function, we receive the equation: sin(x̃) ∗ cos(ỹ) + sin(ỹ) ∗ cos(z̃) + sin(z̃) ∗ cos(x̃) = 0, (2) where x̃, ỹ and z̃ are scaled spatial ordinates, such that x̃ = 2π ∗x/a, ỹ = 2π ∗y/a, z = 2π ∗ z/a. the variable a is the cubic unit cell edge length [4]. we designed 4 types of gyroid structures in total. specimens of all the gyroid structures had same dimensions as the trabecular structures – 14×14×14 mm. the pore size was optimized based on the literature. for example, the optimal pore size for bone ingrowth is 600 µm according to taniguchi et al. [5]. hulbert et al. [6] mentioned that the pore size cannot be smaller than 100 µm for successful bone ingrowth. based on the available literature, the range of 350–800 µm was considered as applicable. gyroid structures number 1, 2 and 3 fulfilled the criterion for bone ingrowth according to the literature, gyroid structure number 4 was slightly out of range. the pore sizes of the designed trabecular and gyroid structures are summarized in table 1. 3. mechanical testing the trabecular and gyroid structures which were produced by additive manufacturing underwent mechanical testing. the processes of manufacturing and mechanical testing were similar for both types of the structures. the goal was to obtain material characteristics of the structures and compare them with the results of the numerical simulation, following numerical model verification. the essential material characteristic for this purpose was the global modulus of elasticity (or global elastic modulus) of each structure. mechanical testing was performed in the form of compression tests. for this purpose, at least three specimens of each structure were produced, 21 specimens of the trabecular structures and 12 specimens of the gyroid structures in total. tests were executed according to international standard iso 13314:2011 mechanical testing of metals – ductility testing – compression test for porous and cellular metals. the standard is intended for porous and foam metals with porosity higher than 50 %. the compression tests were performed on the machine mts alliance rt-30 and rt-50 (mts, usa), which the faculty of civil engineering ctu prague possesses. the maximal load of 30 kn for the trabecular structures and 50 kn for the gyroid structures was 127 p. vakrčka, a. jíra, p. hájková acta polytechnica ctu proceedings a) b) figure 2. the stress-strain curve which is part of the standard for compression testing of porous and foam metals (iso 13314:2011) with specific characteristics, such as 1 – quasi-static gradient, 2 – elastic gradient, 3 – the first maximum compressive strength (σf irst,max), 4 – the compressive offset stress (σ0.2). used. the tests were executed in the laboratory at room temperature. the load direction was the same as the direction of additive manufacturing and thus perpendicular to the printed layers. this was important because of orthotropic behavior, which is caused by the technological procedure. the specimens were loaded by remote displacement and the velocity of deformation was set to 1 mm/min which corresponds to the standard. material characteristics of the trabecular and gyroid structures were based on stress-strain curves which were obtained in the compression tests. the stress was calculated from the following equation: σ = f a , (3) where σ stands for stress, the a stands for the area of specimen cross-section and f stands for the force which is applied perpendicularly to the area of specimen cross-section. the strain (or relative deformation) of the specimen was calculated according to the following equation: � = ∆h h , (4) where � stands for strain, ∆h stands for measured deformation and h is the original height of the specimen. as we mentioned earlier, the most important mechanical characteristic for our purpose was the global modulus of elasticity (e). then we also determined the first maximum compressive strength (σf irst,max) and compressive offset stress (σ0.2). the first maximum compressive strength was set as the first local maximum of the stress-strain curve as can be seen in figure 2a. the compressive offset stress is defined as the stress in material when the material reaches the plastic compressive strain equal to 0.2 % (figure 2b). the quasi-static gradient specifies this moment. furthermore, the angle of inclination of quasistatic gradient was used to determinate the global elastic modulus. although the quasi-static gradient is not equal to the global elastic modulus according to the standard, the curve inclination in this point is the most suitable for the global elastic modulus. at the beginning, the stress-strain curve does not correspond with the real loading progress. behind the point of inflection, the inclination is significantly influenced by the plastic deformation. in more details, we can see the determination of quasi-elastic gradient in ffigure 2b. according to the figure, it can be stated that the quasi-static gradient is defined as the tangent line inclination of the stress-strain curve in the point of inflection. this point divides the curve into the convex and concave parts. the curve concavity is due to plastic deformation. numerically, the point of inflection is specified as the point where the curve inclination starts lowering. it can be calculated from the equation: ∆e = σi �i − σi−5 �i−5 , (5) where σi stands for the stress in the i-step of measurement, �i is the strain in the i-step of measurement, σi−5 stands for the i-5-step of measurement, and finally the �i−5 is the strain in the i-5-step of measurement. the inclination change was set after 5 steps to eliminate the error possibility caused by the measurement deviation of the applied force (f) and deformation (∆h). the next step in determination of the compressive offset stress is detection of the point which is the intersection of the quasi-static gradient and the x-axis (� = 0). numerically, the point can be computed: x0 = − b a , (6) where a,b stands for constants in linear equation of quasi-static gradient in the slope-intercept form. 128 vol. 26/2020 title when the plastic deformation 0.2 % is added (� = 0.2), the intersection of the line with the stress-strain curve gives us the compressive offset stress. the process can be seen again in figure 2b. the slope-intercept equation of the new line is define as: y = a∗x + c c = (x0 + 0.2) ∗a, (7) where a still stands for constant of quasi-static gradient and also for the global elastic modulus of the relevant structure. numerically, the intersection of the mentioned line with the stress-strain curve is the point where following condition is fulfilled: |a∗ �i + c| = |σi| , (8) where σi and �i stands for stress and strain in the istep of measurement. in addition, the last but not less important material characteristic is the porosity of the structures. porosity was determined as: n = 1 − m− (vhom ∗ρt i6al4v ) a∗h∗ρt i6al4v , (9) where m is the weight of the specimen, vhom stands for volume of the homogenous part of specimens, a and h are the dimensions of the porous structures and ρt i6al4v stands for a density of titanium powder which was used to manufacture the specimens. the porosity was used for comparison between the trabecular and gyroid structures. it is assumed that the higher porosity would correspond to the lower global modulus of elasticity. the results of mechanical testing (the compression tests) for each structure are listed in table 1. there are mean values for all sets of the trabecular and gyroid structures. based on the experiment, we can definitely conclude that the global elastic modulus of all the structures is significantly lower than the elastic modulus of the structure material ti6al4v, which is equal to 115 gpa according to the catalogue list. furthermore, the global elastic modulus of all the structures is close to the elastic modulus of human cancellous bone, which is approximately 2.71–9.10 gpa [7]. the crucial part of the design was to adapt global elastic modulus of the structures to the modulus of bone as much as possible to avoid negative consequences which are called stress shield effect. thus, the design of structures met the requirement. the global elastic modulus of trabecular structures is dependent on porosity as we expected. the greatest elastic modulus (3.822 gpa) was showed by the structure with the lowest porosity (0.446) which was rhombic dodecahedron (1.25 mm). on the other hand, the structure with the greatest porosity (0.653), which was rhombic dodecahedron (1.00 mm), had the lowest elastic modulus (2.631 gpa). however, we were surprised that a small difference in the porosity of structure diamond (1.00 mm) and diamond (0.75 mm) should have caused the difference of 18 % in the elastic modulus. the same effect can be seen between structures dode thick (1.00 mm) and diamond (0.75 mm). it is probably caused by the fractions of struts which are discussed at the end of the section. the comparison of trabecular and gyroid structures did not report the same phenomenon. gyroid structure nb 1 with the lowest porosity (0.413) showed significantly lower global elastic modulus (3.048 gpa) than trabecular structure rhombic dodecahedron (3.822 gpa) with similar porosity (0.446). more generally, the gyroid structure showed lower global elastic modulus than the gyroid structures with the same porosity. it is again important to mention that elastic modulus of both types of structures (trabecular and gyroid) was close to the elastic modulus of cancellous bone – 2.71–9.1 gpa. we can conclude that the differences in elastic modulus are not a suitable property for comparison of reliability of the structures. however, the first maximum compressive strength and the compressive offset stress showed huge differences across the structures. it is also important to mention that the compression tests of trabecular structures were executed by using a machine with a lower maximal force of 30 kn. this is the reason why the strength of dode thick (1.00 mm) and rhombic dodecahedron (1.25 mm) were not able to be obtained. although some of the values are unknown, we can state that, in general, the gyroid structures feature higher strength. this fact is crucial because strength is related to the durability of the implant. for our purposes, the global elastic modulus is mainly important in relation to the stress shield effect. on the contrary, the strength is related to the loading which implant can withstand before failure. in the figure 3, we can see the stress-strain curves of all the structures. for clarity, we chose just one characteristic strain-stress curve for each type of the structures. based on the figure, we can see differences in the strength of structures. the trabecular structures have much lower strength than the gyroid structures. what is more, there is a difference at the beginning of the curves. if we take a closer look, we can see that the gyroid structures are linear from the beginning. on the contrary, the trabecular structures have a non-linear beginning. it is assumed that this behavior was due to fractures which were initiated during the fabrication of trabecular structures. it took some time until the material “settled”. the fractures and its origin were described in details in our last paper [8]. in conclusion, we can state that the differences in the elastic modulus of trabecular and gyroid structures are negligible for our purpose. the main difference is that the gyroid structures have higher compressive strength than the trabecular structures. thus, the gyroid structures withstand more loading before failure. 129 p. vakrčka, a. jíra, p. hájková acta polytechnica ctu proceedings type of porous structure pore size e σf irst,max σ0.2 n [µm] [gpa] [mpa] [mpa] [-] gyroid structure nb 1 (1.40 mm) 400 3.048 228.501 166.990 0.413 gyroid structure nb 2 (1.80 mm) 450 2.868 214.805 161.554 0.473 gyroid structure nb 3 (2.40 mm) 750 2.837 190.670 157.764 0.496 gyroid structure nb 4 (3.00 mm) 850 2.772 191.434 154.119 0.515 diamond (0.75 mm) 350 2.884 88.649 86.647 0.540 diamond (1.00 mm) 450 3.508 141.924 126.449 0.549 dode thick (1.00 mm) 500 3.713 n/a 142.218 0.547 dode thick (1.25 mm) 630 2.838 98.211 84.694 0.577 rhombic dodecahedron (1.25 mm) 640 3.822 n/a n/a 0.446 rhombic dodecahedron (1.50 mm) 800 2.631 90.201 78.388 0.653 table 1. the pore size and mean values of the mechanical properties for each type of the trabecular or gyroid structures. figure 3. the overview of the stress-strain curves of trabecular and gyroid structures. we can see quite similar elastic modulus but significantly different compressive strength. 130 vol. 26/2020 title 4. numerical modelling during the compression tests and the microscopical analysis, we evaluated the gyroid structures as more reliable structures than the trabecular ones. it was mainly because of the fractures of struts in the trabecular structures, and because of greater strength of the gyroid structures. for these reasons, we focused only on processing of numerical model of the gyroid structures. the main problem was the process of geometrical modeling and meshing of the structures. it took lot of time to successfully model an appropriate structure without singularities in any modelling software. the most suitable way to achieve the structure without problems was directly in netfabb software by autodesk company. the huge constrain was the limit of ansys academic license. after taking into account the constrain, only element size of gyroid structures was modeled to represent the whole structure. according to gyroid structure number 1, the dimension of cutting box was set to 3 mm. all models were meshed successfully by the automatic meshing method. we had to lower the accuracy of gyroid plotting for structures number 3 and 4 because of the academic license constrain. loading program was set according to iso 13314:2011 mechanical testing of metals – ductility testing – compression test for porous and cellular metals which specify a test method for compressive properties of porous and cellular metals with a porosity of 50 % or more. therefore, all the specimens had to be loaded by the remote displacement 1 mm/min. in general, the rigid body has 6 degrees of freedom which we need to constrain. in consonance with the experiment, we fixed all the degrees on one side of the gyroid structure by fixed support. on the other side, we set the loading by remote displacement. the remote displacement loading had rigid behaviour and ramped rotation around all axes in order to represent the experiment as much as possible. we assumed just elastic part of the stress-strain curve because of implant safety reasons. we do not accept the state when there is plastic deformation, which can lead to the loss of implant. during observation of stress-strain curves of the gyroid structures, we noticed that 150 mpa is the point when all the gyroid structures start the process of hardening. naturally, this value of stress is achieved under different forces based on the gyroid morphology and different strain. now, we would like to discuss the computation of model input for one of the gyroid structures in details. first of all, we had to state the force reaction which is caused by remote displacement. we obtained the value of force which caused stress equal to 150 mpa figure 4. the deformation of the gyroid structure. the unit of deformation is meter. from the experiment data: fmodel gyroid = fexperimental gyroid ∗amodel gyroid aexperimental gyroid fmodel gyroid = fexperimental gyroid ∗ 3 ∗ 3 14 ∗ 14 , (10) where fmodel gyroid stands for force reaction caused by remote displacement, fexperimental gyroid stands for value of force obtained from the experiment which caused stress equal to 150 mpa, amodel gyroid stands for contact loading area of the gyroid which underwent the experimental testing, amodel gyroid stands for contact loading area of gyroid model. the calculated force had to correspond with the force probe on the top of the loaded model. the next step was to determine the final deformation of model. for our purposes, a simplification that strain is equal throughout the structure had to be made. afterward, we could compute the required deformation of our model: umodel gyroid = uexperimental gyroid ∗hmodel gyroid hexperimental gyroid umodel gyroid = uexperimental gyroid ∗ 3 14 . (11) we can see the deformation of gyroid structure in figure 4. in the table 2 we can find the results of curvefitting approximations. we can see that the elastic modulus of material ti6al4v is significantly lower than value according to the material list which is 115 gpa. it is assumed that the reason for such a high difference is due to the problematic loading of samples. the producer of gyroid samples placed a selvage on the edge of platforms. it caused that the loading was distributed mainly through outer parts of gyroid structures, whereas the loading of the numerical model was placed on the complete platform. 131 p. vakrčka, a. jíra, p. hájková acta polytechnica ctu proceedings type of gyroid structure force probe strain of model elastic modulus of material [n] [%] [gpa] gyroid structure nb 1 1350 5.10 14.57 gyroid structure nb 2 1350 5.80 12.98 gyroid structure nb 3 1350 6.10 12.20 gyroid structure nb 4 1350 6.25 11.11 table 2. mean values of the mechanical properties for each of the designed gyroid structures. 5. conclusions the goal of the presented study was to design an effective porous structure which can be used for different types of implants on the interface between bone and artificial material of implant. it can substitute contemporary compact implants with necessary surface modification, and subsequently help to avoid the stress shielding effect and maximize the bone ingrowth. in order to achieve the goal of the study, the different types of trabecular and gyroid structures were designed. both types of the structures were tested macromechanically and their quality was checked microscopically. all the structures showed global elastic modulus close to a cancellous bone. on the other hand, the gyroid structures had greater strength. the gyroid structures, which also seemed to be more reliable based on microscopical analysis, were chosen for designing of the numerical model in the fem software ansys. the results of mechanical testing were also used for the verification of numerical model. all in all, the gyroid structures seem to be an appropriate structure for external layer of implants. as we mentioned earlier, it can speed up the production of implants, which are composed of compact body and porous external layer and do not need any surface modifications. because time and money are always tightly linked, it can also make the implants cheaper. hopefully in the future implants, such as dental, will become a standard and not a nightmare. acknowledgements the financial support provided by the technology agency of the czech republic (tačr), project n. tj01000328, is gratefully acknowledged. the financial support by the faculty of civil engineering, czech technical university in prague (sgs project no. sgs17/168/ohk1/3t/11) is gratefully acknowledged. references [1] t. vieru. mathematics links structure to function in leaves, [online] available at:. https://news.softpedia.com. 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[7] d. wu, p. isaksson, s. j. ferguson, c. persson. young’s modulus of trabecular bone at the tissue level. acta biomaterialia 78:1–12, 2018. doi:10.1016/j.actbio.2018.08.001. [8] a. jíra, p. hájková. trabecular structures as efficient surface of dental implants 2019. 132 https://doi.org/10.1021/acsphotonics.6b00400 https://doi.org/10.1002/adom.201400333 https://doi.org/10.1016/j.msec.2015.10.069 https://doi.org/10.1016/j.actbio.2018.08.001 acta polytechnica ctu proceedings 26:127–133, 2020 1 introduction 2 geometrical model 3 mechanical testing 4 numerical modelling 5 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0546 acta polytechnica ctu proceedings 33:546–551, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague carbon dioxide binding ability in concretes: methodology and modeling lucie schmitta, jena jeongb, jean-marc potierc, laurent izoretd, jonathan mai-nhua, ∗, nicolas decoussera, thomas pernina a cerib: centre d’etudes et de recherche de l’industrie du béton, 1 rue des longs reages, 28230 épernon, france b estp: école spéciale des travaux publics, 2-4 rue charras, 75009 paris, france c snbpe: le syndicat national du bêton prêt à l’emploi, 32 allées d’orléans, 33000 bordeaux, france d atilh: association technique de l’industrie des liants hydrauliques, la defense, 7 place de la défense, 92974 paris-la-défense cedex, france ∗ corresponding author: j.mai-nhu@cerib.com abstract. carbonation of concretes is a natural physico-chemical process that can be described as a reaction between the carbon dioxide contained in the air and the cement matrix. carbonation concerns all concretes types in contact with the ambient air but also concretes in ground, from production stage to use and end-of-life stages. the amount of carbon dioxide bound varies according to the type of binder, the compacity of concrete and the environmental conditions during the use and the end-of-life stages. to consider the re-carbonation of concrete, works have been carried out within the framework of the european standardization group cen/tc229/wg5 and in cen/tc104. a methodology to consider the re-carbonation of concrete structures has been proposed in the nf en 16757 standard on environmental product declarations for concrete and concrete elements. in addition, fd cen/tr 17310 provides detailed recommendations regarding carbonation and absorption of carbon dioxide in concrete and give some precisions for application of nf en 16757. this is an important topic towards a sustainable development in the current context of circular economy and co2 uptake related to the french energy labelling (e+c-). in this paper, numerical and analytical carbonation models are used to estimate the co2 binding ability of concrete structures. the obtained results are compared to the methodology proposed in appendix bb of nf en 16757 standard. they confirm that the methodology described in the nf en 16757 standard leads to estimated degree of carbonation of the same order of magnitude. the advantage of using more advanced models lies in a better consideration of environmental parameters, the possibility to simulate the behaviour of crushed concrete, its reuse in new concrete as recycled aggregate and the possibility to simulate the carbonation of concretes in ground. this is an immediate perspective in the ongoing work in the french national project fastcarb on accelerated carbonation of recycled concrete aggregates. keywords: concrete products, concrete structures, modeling, re-carbonation, standardization. 1. introduction concrete is the most widely used building material in the world because the raw materials are generally considered lowly exhaustible and ubiquitous on the globe. in addition, its manufacturing process is well controlled and its cost price low. however, due to the use of cement and its success in construction works, concrete production can generate considerable co2 emissions. today the cement sector is among the main emitters with between 5 and 7% of global emissions [1]. how can this situation be reconciled with, on the one hand, the need for countries to continue to offer an efficient, robust and affordable solution and, on the other hand, the need to fight effectively against global warming? the building sector is actively working on solutions to reduce its co2 emissions. current research considers many aspects: • to reduce the carbon intensity of cement production operations; • to control the environmental impact of concrete products on their entire life cycle; • to optimize the processes of concrete production and construction; • to better recycle concrete into concrete; • to work in partnership with the construction industry to jointly contribute to green building systems. this study is part of a national research project: the fastcarb project. the aim of the fastcarb project is firstly to store co2 in the recycled con546 https://doi.org/10.14311/app.2022.33.0546 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 binding ability of carbon dioxide in concretes figure 1. carbon footprint of 1kg of clinker. type of cement carbon footprint* cem i 765 kg of co2/ton cem ii/a-s cem ii/a-l 671 676 kg of co2/ton cem ii/b-l or ll cem/b-m 579 585 kg of co2/ton cem iii/a 400 kg of co2/ton cem iii/b 274 kg of co2/ton cem v/a 468 kg of co2/ton * not including carbon dioxide emissions from the combustion of secondary fuels table 1. carbon footprint of the different types of cement in france (in average) [2]. crete aggregates (rca), in order to reduce the environmental impact of concrete in the structures and secondly to improve the quality of these aggregates by reducing their porosity. today one of the main issues for the concrete profession is to better evaluate the carbon footprint of concrete and concrete structures. to solve this question, the quantity of carbon dioxide that could be stored by a building during its life time has to be correctly calculated. the objective of this study is to quantify the co2 stored by a building during its life time using different methods and models. 2. context and objectives of the study cement is a widely used building material whose main hydraulic constituent is clinker. to manufacture the clinker, the active "raw material" of the cement, the calcination of the limestone and the clay in the kilns at very high temperature (1,450 ◦)c generates carbon dioxide due to decarbonation of limestone and to combustion reactions during pyroprocessing. this decarbonation step represents about 60% of co2 emissions of the production process [1, 3, 4]. the figure 1 presents the carbon footprint of the production of clinker and the table 1 gives the footprint of the different types of cement. the carbon dioxide emitted during the manufacturing of cement can for a part be stored again by the natural phenomena of carbonation during the concrete life time. during the life of the structure, the concrete could store carbon dioxide at a level of 10 to 15% of the co2 emitted during the decarbonation of the limestone necessary for the manufacture of cement [1]. it is important to note that this percentage is average on the scale of the building. some parts carbonate more than others (exposed/not exposed parts, porous masonry compared to civil engineering parts. . .). these numbers take into account decarbonation and combustion of primary fuels but do not consider co2 emissions due to the combustion of "wastes" like tyres or waste oils for example. this is this value which is used for the lca. the standard nf en 16757 presents a methodology and a model to evaluate the quantity of carbon dioxide that can be stored during concrete lifetime in function of the environment and the materials as concrete strength classes. in this study, the numerical model sdream-crete is used to estimate the capacity of co2 storage of concrete. the results are compared to those obtained with the nf en 16757 model. then the co2 storage capacity of a r+5 building is estimated and compared to the amount of carbon dioxide emitted during its construction. 3. presentation of the models 3.1. model of the standard nf en 16757 [5] the nf en 16757 standards model is an analytic model which considers environment as well as material parameters. the equation 1 gives the carbonation depth. the equation 2 presents the model to 547 l. schmitt, j. jeong, j. m. potier et al. acta polytechnica ctu proceedings material environment xc: carbonation depth (mm) (equation 1) x x k: carbonation rate (mm/year0.5) x x kk: correction to k−factor x − t: time (year) − − utcc: maximum theoretical uptake in co2/kg of cement (0.49 for cem i) x − c: cement content (kg/m3) x − dc: degree of carbonation (−) − x table 2. parameter of the nf en 16757 model. figure 2. schematic presentation of the sdream-crete model (only carbonation and hydric parts). properties value considered resistance class c25/30 cem i 52.5 n [kg/m3] 280 aggregates [kg/m3] 1 922 efficient water [l/m3] 168 wef f /c 0.60 relative humidity [%] 80 ± 10 co2 concentration [%] 0.04 table 3. properties of studied concrete and environment. estimate the co2 stored and the table 2 presents the parameters. xc = k · √ t (1) co2 uptake = k · kk ! √ t 1000 " utcc · c · dc (2) each of these factors is detailed in appendix bb of the standard. the appendix bb also gives values for the parameters that depend on the material and the environment. 3.2. model sdream-crete the sdream-crete durability model was developed at cerib with lmdc [5, 6]. it is a numerical model based on mass conservation equations. this model can simulate the chloride ingress, the carbonation, the moisture transfers as well as the development of the corrosion of the rebars in concretes exposed to wetting-drying cycles. it can be used in a deterministic context or in a probabilistic context to achieve reliability calculations for concrete products or concrete works. in this study, it is used to calculate the carbonation depth and then the amount of co2 absorbed by the concrete during its service phase. the quantity of stored co2 is calculated with the carbonation depth and the trapezoidal rules. the figure 2 presents a schematic representation of the sdreamcrete durability model. 4. results 4.1. comparison between analytical method and numerical method the first phase of this study consisted of estimating the amount of co2 that can be absorbed per square meter of exposed concrete with both models presented in the precedent section. the properties of the considered concrete and environment are presented in the table 3. 548 vol. 33/2022 binding ability of carbon dioxide in concretes properties value considered k: carbonation rate (mm/year0.5) 1.6 kk: correction to k-factor 1 (cem i) utcc: maximum theoretical uptake in co2/kg of cement 0.49 (cem i) dc: degree of carbonation (-) 0.85 (outdoor exposed to the rain) table 4. details of the considered value for the calculation with the nf en 16757 model. figure 3. carbonation depth of the studied concrete calculated with sdream-crete model and nf en 16757 model. figure 4. absorbed carbon dioxide calculated with two models and two methods. the figure 3 presents the carbonation depth calculated with sdream-crete model and with the equation 1 (en 16757 on the figure). the two models (sdream-crete et eq. 1) provide results slightly different. that can be explained by the different parameters taken into account by the two models. the standard model is based on the law at the root of time while the numerical model considers more precisely the formulation and the environment of the concrete. that will be discussed later. considering that the carbonation reaction can be summarized as (eq. 3): ca(oh)2 + co2 ↔ caco3 + h2o (3) the equation 3 means that the consumption of one mol of co2 leads to the production of one mol of calcite. with sdream-crete model, the amount of calcite in the concrete can be calculated and, by extension, the amount of carbon dioxide consumed according to the equation 3. the figure 4 presents the quantity of stored co2 for one square meter of the considered concrete. the orange curve was obtained using the equation 2. the blue curve was obtained using the depth of carbonation calculated by sdream-crete and then using the mathematical method of the trapezoids. these results show that two different models and methods, that consider very different parameters, are leading to very close results in this specific scenario. both method and model have advantages and disadvantages: the analytical model has the advantage of being quick and easy to use. in fact, the standard nf en 16757 details each of the parameters and proposes values to be considered depending on the resistance of the concrete and its environment. the numerical model seems more complete than the standard model because of the phenomena it considers (water cycle, detailed composition of concrete...). however, 549 l. schmitt, j. jeong, j. m. potier et al. acta polytechnica ctu proceedings figure 5. illustration of the building modelized for the study. outdoor indoor k: carbonation rate (mm/year0.5) 1.6 (exposed to rain) 4.6 (with cover) kk: correction to k-factor 1.05 1.05 t: time (year) 50 50 utcc: maximum theoretical uptake in co2/kg of cement 0.41 0.41 c: cement content (kg/m3) 300 300 dc: degree of carbonation (−) 0.85 (exposed to rain) 0.4 (dry climate) quantity of stored co2 (according to eq. 2, kg/m2) 1.24 1.68 table 5. values considered for the study. the calculation times can be longer and obtaining the input data can sometimes be complicated in an operational context. in this study, the standard model will be used (eq. 2). 4.2. application to a five-storey building [7] in this paragraph, the nf en 16757 model is applied to a r+5 building made with cem ii/a (case of an exposed architectural concrete, we only take into account the facade, external and internal) and the result is compared to the amount of co2 emitted during the decarbonation of the clinker needed for the building construction. the figure 5 presents an illustration the fivestorey building studied. the first step of the study is to determine the quantity of carbon dioxide emitted during the building construction (only considering decarbonation of cement). the cement used for the construction of this building is cem ii/a-l and contains 80% clinker. the amount of clinker required for this building is therefore 48 384kg. then it is possible to calculate the quantity of carbon dioxide emitted during the decarbonation phase (0.54kg for 1kg of clinker): 26ă127kg of co2 emitted. then we can calculate the quantity of carbon dioxide stored by the building with the model of the nf en 16757 standard (eq. 2). the table 5 presents the values of the parameters used for the simulation. all these values were determined using annex bb of standard nf en 16757. after calculating the amount of carbon dioxide that can be absorbed per square meter of concrete, we can calculate the amount of co2 that can be stored by this building after 50 years of life (eq. 4). total co2 uptake =# outdoor co2 uptake + indoor co2 uptake $ · surface of exposed concrete (4) so, the quantity of carbon dioxide that can be stored by the building is 2ă943.4 kg co2. if we compare this value to the quantity of co2 emitted during the construction (by the decarbonation phase, 26 127kg) we can highlight that the ratio is around 11.3%, which corresponds to the value of the literature [1] (eq. 5). total co2 uptake emitted co2 decarbonation = 2943.4 26 127 = 0.113 (5) this study was carried out on a fictitious and heavily simplified building at first. a single concrete for the whole building has been considered but in reality, we can observe significant divergences due to the possible variety of concretes encountered, e.g., porous masonry concrete or more efficient and dense concretes in some cases. the direct following will be to do the same work on an existing building for which we have all the material and environmental data. both models can be used. 550 vol. 33/2022 binding ability of carbon dioxide in concretes 5. conclusion as an integral part of the national project fastcarb, this study made it possible to carry out a bibliographic study on the ability of concrete to capture co2 during its service phase. several models of different levels (analytical, numerical ...) exist today. nf en 16757 is the product category rules for concrete and concrete elements. one model to take into account the carbonation phenomena in the context environmental product declaration (epd) for the calculation methodology is presented in the nf en 16757 standard. this standard is the reference document when performing epds for concrete elements following en 15804+a1 standard. moreover fd cen/tr 17310 [8] provides technical and scientific elements to support the carbonation treatment part of the nf en 16757. in the first part of this study, two different models and methods for calculating the quantity of stored co2 were compared. the results show that the model of standard nf en 16757, although analytical and "simple", allows to obtain results close to those obtained with a numerical model taking into account many different parameters. then, the model proposed in the standard was used to assess the amount of carbon dioxide captured by a five-storey building during its service phase (only the facade). the calculations highlight the ability of the concrete to capture part of the co2 emitted during the decarbonation of the clinker. with approximately 11.3% of the emitted co2 which can be reabsorbed, the results confirm the literature. it is important to note that this value could be increased significantly if carbonation continued after the demolition of the structure. this last point is the direct perspective of the work reported in this article. recent results show that it is possible to recapture up to 50 to 60% of the co2 emitted during the decarbonation of limestone for traditional concrete [1]. this value including the percentage of carbonation during the service life of the building, this means that the recycling phase alone stores around 40% to 45% of the co2 emitted during the decarbonation of limestone. investigations and calculations are currently underway to precise the co2 storage capacity of a recycled concrete aggregate (rca) and the best condition to optimize the carbonation of rca. acknowledgements the research and results reported herein are supported by the french ministry in charge of construction under the fastcarb research program. references [1] n. roussel. carbonatation des bétons et piégeage du co2, solution béton (hors série), p. 1-8, 2012. [2] atilh. ciments : déclaration environnementale, inventaire & analyse du cycle de vie, 2018. [3] r. andersson, h. stripple, t. gustafsson, et al. carbonation as a method to improve climate performance for cement based material. cement and concrete research 124, 2019. https: //doi.org/10.1016/j.cemconres.2019.105819. [4] gcca. gnr global database on co2 and energy information. https://gccassociation.org/. [5] schmitt l, mai-nhu j, duprat f, de larrard t, rougeau p, 2019, interest of using a model combining carbonation/chloride ingress and depassivation to better anticipate the rehabilitation of concrete structures, international conference on concrete repair, rehabilitation, and retrofitting, cape town, south africa. [6] j. mai-nhu, p. rougeau, f. duprat, et al. probabilistic optimization of the concrete cover for structures exposed to chlorination and carbonation according to climate change. safety, reliability, risk and life-cycle performance of structures and infrastructures 199:1535-42. 2013. https://doi.org/10.1201/b16387-225. [7] l. schmitt, j. jeong, j. m. potier, izoret l, et al. carbon dioxide binding ability in concretes: methodology and modeling, international workshop co2 storage in concrete co2sto2019 france marne la vallée ifsttar, 2019. [8] afnor. carbonation and co2 uptake in concrete, fd cen/tr 17310, 2019. 551 https://doi.org/10.1016/j.cemconres.2019.105819 https://gccassociation.org/ https://doi.org/10.1201/b16387-225 doi:10.14311/app.2021.30.0018 acta polytechnica ctu proceedings 30:18–23, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app finite element implementation of geometrically nonlinear contact ondřej faltus∗, martin horák czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: ondrej.faltus@fsv.cvut.cz abstract. the oofem finite element software has been recently updated to include contact algorithms for small strain applications. in this work, we attempt to extend the contact algorithms to large strain problems. reviewing the current code and comparing it with approaches encountered in literature, we arrive at a specific algorithmic solution and integrate it into the current code base. the current code is explained, the necessary extensions are derived and documented, and the algorithmic changes are described. tests confirm the functionality and quadratic rate of convergence of the proposed implementation. keywords: contact, large strains, node-to-segment, oofem, penalty method. 1. introduction having originated with heinrich hertz in 1881 [1], contact mechanics has been a fringe discipline of structural mechanics for most of the 20th century, owing mainly to its limited options for analytical analysis. only with the advent of numerical methods, such as the finite element method (fem), has development in this field sprung forward [2]. this surge in new advances is only exacerbated by the ever more rapid evolution of computational hardware capabilities experienced today; ever more complicated contact problems necessitating ever more complicated methods previously unimaginable due to hardware constraints are coming into focus [2, 3]. the tool used for the purposes of this article is the finite element computational software oofem [4]. oofem is a free, open-source c++ code developed mainly at the czech technical university in prague. it is based on object-oriented design with emphasis on variability and extendability. currently, oofem’s finite element code supports a number of modules for solution of various types of numerical problems. of interest for our purposes is the sm module intended for structural mechanics. basic introduction of contact algorithms into the oofem software has been achieved in [5]. this implementation, however, has certain limitations, among them chiefly the lack of support for large strain, i.e., geometrically non-linear problems. in this work, we strive to remove this particular limitation and demonstrate a simple method of extending existing basic penalty-based formulation to handle large strain problems. 2. algorithm 2.1. basic equation without loss of generality, let us consider two deformable bodies bγ , γ = 1, 2. moreover, let ωγ0 and ωγ represent the regions occupied by the body in the reference and deformed configuration, respectively. position of a point in the reference configuration is denoted by x γ , and a position xγ of a point in the deformed configuration is obtained be means of the one-to one deformation mapping ϕγ as xγ = ϕγ (x γ ) (1) the basic equations for each body under consideration are the following: • kinematic equations f γ (x γ ) = ∂ϕγ (x γ ) ∂x γ ∀x γ ∈ ωγ (2) • equilibrium equations divp γ (x γ ) + b̄γ (x γ ) = 0 ∀x γ ∈ ωγ (3) • dirichlet boundary conditions uγ (x γ ) = uγ (x γ ) ∀x γ ∈ γγu (4) • neumann boundary conditions t γ (x γ ) = p γ (x γ ) · n̄ γ (x γ ) ∀x γ ∈ γγσ (5) • constitutive equation p γ (f ) = λγ tr(eγ )f γ + 2µγ f γ · eγ (6) here, we use the so-called saint-venant kirchhoff material model. note that the contact implementation is not affected by the choice of constitutive law and extension to inelastic law is straightforward. 18 http://dx.doi.org/10.14311/app.2021.30.0018 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 fem implementation of geometrically nonlinear contact in the equations above, f is the deformation gradient, e is the green-lagrangian strain tensor, div is divergence operator in the reference configuration, p is the first piola-kirchhoff stress, b̄ are prescribed body forces in the reference configuration, u are prescribed displacements, n̄ is a unit normal vector in the reference configuration, γu and γσ are the dirichlet and neumann boundaries which are mutually disjoint, i.e., γγu ∩ γγσ = ∅, and γγu ∪ γγσ = γγ , λγ and µγ are lame’s parameters. moreover, contact conditions introduce essentially a new set of boundary conditions into the task. note that only normal contact is considered in what follows. the core issue of contact problems can be described by means of the contact constraints in the form of karush-kuhn-tucker optimality conditions [6]: tc ≤ 0 gc ≥ 0 gctc = 0 (7) the first condition represents the fact that contact forces tc are only compressive, the second condition is characterized by a gap function g(u) and express a non-penetration condition. the last equation is simply complementarity between two distinct and mutually exclusive states for the system; either contact occurs, and therefore the gap function gc is zero, while a non-zero compressive contact force is present, or the contact force is zero and the gap function is larger than zero. 2.2. penalty method due to the constraint conditions in form of ineqaulities, see (7), weak form of the above introduced formulation is represented by a variational inequality, see [2]. several methods how to handle the contact constraints are possible, e.g., the lagrangian multiplier method, the augmented lagrangian method, or the nitsche method [3, 6]. currently in oofem, the lagrangian multiplier method and the penalty method are implemented [5]. it is the penalty method which we have selected for the purpose of large strain extension, as its formulation is more straightforward and the necessary code complexity is notably lower. for the penalty method, the weak form can be written as ! γ " ωγ δf γ : p γ − δuγ · b̄γ dv − − " γγσ δuγ · t̄γ dγ + w c = 0 (8) where δ represents variation and w c is the penaltybased contact contribution along contact interface γc defined as w c(u) = " γc 1 2 h(−g(u))pg2c (u) dγ (9) where h is the heaviside function and p is the penalty parameter. the main disadvantages of the penalty method is the that the achieved solution is imprecise (it would only be precise for p → ∞), but moreover, any improvement in precision acts directly against the stability of the computation, as larger penalty parameters introduce larger members into computational stiffness matrices, which disturbs most numerical solvers [2]. those disadvantages are balanced by the relatively simple formulation of the method and, in comparison with the lagrangian multiplier method, by simple resulting equation systems, as no new variables are introduced into the computation. 2.3. node-to-segment discretization an implementation of contact problems into the framework of the finite element method necessitates selecting a distinctive approach at contact discretization. as the physical world is described by fem in the form of finite elements and connecting nodes, contact can only be considered between those. typical contact discretizations include nodeto-node, node-to-segment, segment-to-segment and other approaches [3]. at current time, node-to-node and node-to-segment discretizations are implemented in oofem [5]. we are only extending the node-tosegment discretization for large strains, as the nodeto-node discretization is by definition inadequate for this purpose. in node-to-segment discretization, a single node comes into contact with a segment. various objects may be understood under the term "segment", most prominently a set of element boundaries (element edges in 2d domains or element surfaces in 3d domains) or an arbitrarily defined rigid analytical surface. contact detection in node-to-segment discretization necessitates determining three things [3, 6]: (1.) the gap function gc, essentially a projection of the node onto the contact segment, which determines whether there is contact or not; analogical to the penetration function in the analytical approach to contact problems. (2.) the normal vector in the deformed configuration n, which determines the orientation of the contact surface at the contact point. obtained for most segments in oofem as vector perpendicular to the contact segment’s surface at the contact point [5]. (3.) the extended n-matrix n ∗, which ensures allocation of the external forces and stiffness to the proper degrees of freedom in the fem formulation. when the segment in question is an element boundary, the matrix takes the form [5] n ∗ = # i(d×d) −ne(xc) $ (10) where i(d×d) is a square unit matrix of the appropriate dimension, which corresponds to degrees of freedom of the contacting node, and ne(xc) is the 19 ondřej faltus, martin horák acta polytechnica ctu proceedings matrix of element boundary interpolation functions evaluated at the contact point. 2.4. fem formulation to integrate contact constraints into the fem framework, considerations have to be made for the contact forces and energy discussed above within the principles of fem discretization. for the penalty method, contact energy is expressed as w c = 1 2 pg2c (11) note that the energy contact contribution above is calculated only in the contact point, which can be interpreted as lobatto integration of (9), see, e.g., [6] for more details. variation of this part of energy yields δw c = pgcδgc (12) it is worth noting that the term pg(u) represents the contact force. variation of the gap function can be further expressed as δgc = % n ∗t n ||n|| & δdc (13) where δdc are the variational nodal displacements of relevant fem nodes and n is the normal vector in deformed configuration. the external forces of contact are therefore fc = −pgc % n ∗t n ||n|| & (14) to obtain the associated tangential stiffness matrix. the internal forces have to be subjected to differentiation with respect to nodal displacements dc [5]. for small strains, it is generally assumed that the direction of contact does not change throughout the analysis, and therefore n remains independent of dc [5, 6]. kc = − ∂fc ∂dc = p % n ∗t n ||n|| &t % n ∗t n ||n|| & (15) 2.5. extension to large strain problems now we stand before a question: which of the previously described considerations change if we consider large strains? the mechanics of introducing contact constraints to fem remain the same, as well as the definition of contact energy. two critical differences appear. firstly, it can no longer be assumed that contact follows the same direction throughout the whole computation. the normal vector has to be reassessed after each computational step and also the stiffness matrix has to reflect this. secondly, large deformations result in change of contact point on the segment, which is reflected in the evaluation of matrix n ∗. the normal vector is deformation dependent and can be expressed geometrically [2] n = t1 × t2 (16) where × is the vector cross product and ti are tangential vectors of the element boundary in the deformed configuration (in the case of an element edge, i.e., contact formulation in 2d, one is the tangent and the other is the e3, that is, a unit vector perpendicular to the plane). for the new expression of stiffness, we elect to utilize a formula presented in [2]. specifically for the case of node-to-2d-segment penalty method, it reads k = p # nv n t v − gc l # bv t t v + tv btv + gc l bv b t v $$ (17) with nv = n ∗t n ||n|| bv = b∗t n ||n|| tv = n ∗t t (18) where b∗ is the extended b-matrix. note that the first member of the stiffness matrix corresponds with the previously derived term for the small strain case. for a linear edge in 2d, the extended matrices are defined as n ∗ = ' −1 0 ξ − 1 0 −ξ 0 0 −1 0 ξ − 1 0 −ξ ( (19) b∗ = − ∂n ∗ ∂ξ = ' 0 0 −1 0 1 0 0 0 0 −1 0 1 ( (20) where ξ is the parametric coordinate on the element edge. 3. implementation 3.1. class structure the oofem code has a strictly object-oriented structure [4]. the integration of contact into this structure has been mainly accomplished in [5] and only slight additions were necessary for the large strain extension. figure 1 displays the relevant parts of oofem class structure together with the classes used for contact. the general idea of class structure centers around two main classes: elementedgecontactsegment and node2segmentpenaltycontact. elementedgecontactsegment is derived from the general contactsegment class and represents element edges in 2d. when initialized from an oofem input file, the class receives a set of element boundaries. internal processes ensure that contact is considered always with the edge which is closest to the contacting node [5]. node2segmentpenaltycontact is a subclass of activeboundarycondition, an oofem feature 20 vol. 30/2021 fem implementation of geometrically nonlinear contact figure 1. a diagram of oofem class structure, focused on contact-relevant parts, created with the help of [7]. which allows boundary conditions to contribute by themselves to the internal forces, external forces, and stiffness matrix. the utilization of this feature avoids creating any form of a "contact element", otherwise prevalent among fem contact implementations [3]. the contact boundary condition is likewise initialized from the oofem input file with information on the size of the penalty parameter, a set of nodes and a set of one or more contact segments. any node is always checked for contact with any segment [5]. 3.2. operations the assembly of external forces and stiffness matrix takes place in the node2segmentpenaltycontact class, specifically in its overridden methods assemblevector() and assemble(). for the determination of gap gc, the contact segment method computepenetration() is invoked, which computes it from a purely geometrical consideration, considering the node to have penetrated the segment when its deformed coordinates are on a different side of the segment than the undeformed ones. the contact segment is also responsible for the computation of the tangent and normal vectors. the tangent vector is determined as t = ∂x ∂ξ = −b∗seg xn = (21) = ' −1 0 1 0 0 −1 0 1 ( ) **+ **, x1 y1 x2 y2 **. **/ = 0 x2 − x1 y2 − y1 1 where b∗seg is the segment part of the extended bmatrix and xi, yi coordinates of the segment nodes. meanwhile, the normal vector is determined as n = e3 × t (22) for large strains and from the element::edgeevalnormal() function for small strains (which returns the undeformed normal vector). neither of those vectors is normalized. the normalization of the normal vector happens in the contact condition, when the nv , bv and tv matrices are determined according to (18). in the assembly of stiffness, the contact conditions asks the contact segment whether the large strain part shall be assembled. this is accomplished by the contactsegment::hasnonlineargeometry() function and based on the setting of the element in contact. 4. demonstration to demonstrate the functionality of our approach, we select a simple computational example in a 2d plane strain idealization. the mesh used can be seen in figure 2a. a double-clamped beam is positioned above a cantilever. the length of the beam is 5 m and length of the cantilever is 3 m. both are 0.25 m high with a gap of 0.25 m between them. the top of the beam is loaded in every step by an uniform distributed load of 8000 kn/m, causing vertical deformation and thus 21 ondřej faltus, martin horák acta polytechnica ctu proceedings (a) . undeformed mesh. (b) . deformed state and vertical component of stress in step 4 (initiation of contact between beam and cantilever). (c) . deformed state and vertical component of stress in step 200 (end of analysis). figure 2. a computational example demonstrating the code and its convergence. iteration force err. x disp. err. x force err. y disp. err. y 1 5.074e-05 0.000e+00 5.450e-03 0.000e+00 2 7.311e-03 8.044e-02 1.583e-01 4.515e-02 3 1.879e-04 2.920e-03 4.120e-03 2.014e-03 4 1.465e-04 2.844e-03 3.202e-03 1.062e-03 5 2.041e-07 5.071e-04 1.488e-06 1.809e-04 6 1.418e-12 1.863e-06 3.298e-11 7.549e-07 table 1. evolution of global force and displacement errors in iterations of loading step 14. subsequently contact between the two bodies. loading is performed in 200 uniform loading steps. both beam and cantilever are modelled by rectangular plane strain elements (oofem element quad1planestrain). in total, the task mesh consists of 805 elements. refinement of the mesh has been accomplished with the help of a custom script in the matlab software [8]. all elements are made of the same saint-venant kirchhoff material with the parameters λ = 100 mpa and µ = 150 mpa. contact conditions are introduced between the lower edge of the beam and the upper edge of the cantilever below. on the latter interface, two boundary conditions are defined, one pairing the beam’s nodes with the cantilever’s element edges and the other vice versa. this ensures that no clipping caused by crude or otherwise imperfect meshing can take place. figure 2b shows the deformed state of the mesh and the vertical component of the stress tensor in step 4 of 200, when the beam and the cantilever first come into contact. likewise, figure 2c shows the same in step 200 at the end of the analysis. to render those 22 vol. 30/2021 fem implementation of geometrically nonlinear contact figure 3. evolution of error during the equilibrium iteration process. figures, the paraview open-source software was used [9]. values of stress in the visualizations are given in pa. note that between steps 4 and 200, the contact surface moves significantly to the left both on the beam and the cantilever. this was accomplished by alternating activation of the two opposite boundary conditions on the interface, which managed to subsequently introduce contact between neighboring nodesegment pairs, while the contact gap between the original pairs opened up again. the main objective of this test was to prove quadratic convergence of the new approach to large strain contact stiffness matrix. the vast majority of the 200 steps computed managed to converge in 6 iterations or less. the only exception is step 4, in which the contact occurs for the first time. table 1 shows error evolution in a typical computational step (step 14 was chosen). moreover, the dependence of error (norm of residual forces) on the number of iterations during the equilibrium newton-raphson process for steps 4, 14, and 164 is visualized in figure 3. thanks to the semilogarithmic scale used, it is apparent that approximate quadratic convergence has been achieved. 5. conclusions the paper describes an implementation of a two dimensional large strain node-to-segment penaltybased approach to contact mechanics into an opensource finite element code oofem. the emphasis is given on proper design of the hierarchy of classes with a focus on modularity and extensibility towards different methods for handling contact constraints, contact with friction, or three dimensional formulation. the functionality of the presented approach is successfully verified on a numerical example and a quadratic rate of convergence of the newton-raphson iteration procedure is demonstrated. acknowledgements the authors would like to acknowledge support received for work on this project from a ctu grant no. sgs20/038/ohk1/1t/11. references [1] h. hertz. on the contact of elastic solids. z reine angew mathematik 92:156–171, 1881. [2] p. wriggers. computational contact mechanics. springer, new york, 2nd edn., c2006. [3] v. a. yastrebov. numerical methods in contact mechanics. wiley, hoboken, nj, 2013. [4] b. patzák. oofem home page, 2000. http://www.oofem.org. [5] o. faltus. object-oriented design and implementation of contact mechanics into finite element code oofem. master’s thesis, czech technical university in prague, 2020. [6] a. konyukhov, r. izi. introduction to computational contact mechanics. wiley, chichester, west sussex, 2015. [7] staruml. version 3.2.2. mklabs co.,ltd., seoul, republic of korea, 2020. [8] matlab. version 9.3.0 (r2017b). the mathworks inc., natick, massachusetts, 2017. [9] u. ayachit. the paraview guide: a parallel visualization application. kitware, 1st edn., 2015. 23 https://doi.org/10.14311/app.2022.33.0617 acta polytechnica ctu proceedings 33:617–623, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague breakthrough of the resemblances and correspondences between resilience and sustainability in civil infrastructures oscar javier urbina∗, elisabete teixeira, hélder sousa, josé matos university of minho, isise, department of civil engineering, guimars̃es, portugal ∗ corresponding author: oscarj105@gmail.com abstract. sustainable construction has become a growing trend among researchers and stakeholders. simultaneously, resilience and risk assessments for civil infrastructures have flourished in terms of importance among researchers, economic sectors, and society. nevertheless, there is no abundant research that correspond to both approaches, despite that, there are massive similarities and shared characteristics between both investigation branches. distinctively, this year has demonstrated that sustainable development is directly obstructed by different extreme events that trigger risks and vulnerabilities in civil infrastructures. these extreme events require a deep and complex study to minimize the impacts they may cause in society and economy, two main factors considered in the study of sustainability. therefore, when a risk and resilience assessments are conducted, it is already analyzed as a sizable part of sustainability. consequently, there exists a possibility to create a methodology that examines and assesses four categories of civil infrastructure sustainability: technical, environmental, social, and economical. the aim of this paper is to demonstrate the need of a comprehensive approach between sustainability, risk, and resilience assessment, compiling and comparing the existing methodologies for assessing the impacts on civil infrastructures, showing that both present resemblances and none can be omitted, being necessary for the decision-making. keywords: civil critical infrastructures, extreme events, resilience assessment, risk assessment, sustainable construction. 1. introduction world population has increased dramatically in the last decades, with a difference of 6 billion people between only one century. this exponential increment of population results on an excessive use of primary resources in order to fulfill their needs, being this one of the explanations towards the climate change problem we are currently facing [1, 2]. as a result of this worldwide problem, in the late 70s, reports began to appear concluding that if the population growth and its consumption of resources are not controlled, the world would not withstand the pressure on its resources, giving birth to the term of "sustainable development" [3]. nowadays sustainability is a common concept in research that has been the focus of multiple studies, these investigations usually determine the development of humanity in the interconnection of three dimensions, the economic dimension, which is the most widespread over time, the environmental dimension, and the social dimension [3, 4]. also, it is clearer now that the sustainable development is occasionally obstructed by new risks and vulnerabilities, such as terrorist attacks or diseases, as we are currently facing the new sars-cov-2 [5]. as emphasized in the united nations agenda 21, there are some risks, designated as systemic risks, that are arising actively (e.g. diseases, terrorism, natural disasters due to climate change), facing them in a neverending challenge loop [5]. today, critical infrastructure systems face an increased number of hazards, such as: natural (earthquakes, floods, fires), technological (operational failures in systems), and human (fires, cyber-attacks, or terrorism), that can intervene in the functionality of these systems. the malfunction of these systems can cause a cascading effects through the community that produce social, economic, and functional disruption [6, 7]. therefore, there is the need to identify the new upcoming risks that may affect infrastructures, so at the time needed, governments address hazards within a comprehensive management and understanding of the complexity of civil infrastructures. risk management models and resilience assessments are the best solution to assess and find solutions for planning, mitigation, and recovery for civil infrastructures under extreme events, where the consequences of damage and losses are quantified for decision making [7]. in this context, risk and resilience analysis play an important role, as it provides information that help decision makers to develop risk mitigation plans and strategies, taking into account that the economic development of a country strongly depends on its level of infrastructures [8]. in consequence, these concepts should be brought into a comprehensive approach that leads towards a sustainable development 617 https://doi.org/10.14311/app.2022.33.0617 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en o. j. urbina, e. teixeira, h. sousa, j. matos acta polytechnica ctu proceedings of infrastructures [5]. the aim of this paper is to demonstrate the need of a comprehensive approach between sustainability and risk and resilience assessment, compiling and comparing the existing methodologies for assessing the impacts on civil infrastructures, showing that both present resemblances and none can be omitted, being necessary for the decision-making. thus, this paper seeks to establish the theorical foundation for a future approach that combines and evaluates the concepts of resilience, risk, and sustainability of civil infrastructures simultaneously using quantitative and qualitative methods. as this is of high importance for stakeholders, authorities, and industries in the decision making for civil infrastructures, under any hazard and the impacts that it would incur in social, economic, environmental and functionality terms. this paper is mainly developed in 5 chapters. the first chapter consists on a brief introduction, which includes a general framing of the topics, the definition of the objectives of the work and a presentation of the structure of the article. in the second chapter, it is introduced the concept of sustainable development, as it is necessary to understand the context in which it emerged, as well as its historical evolution, then a short display of the most known methods at international level for sustainability in construction are presented, showing their characteristics and some examples of the indicator measures are introduced. in the third chapter it is addressed the beginning of risk and resilience analysis, recalling on some of the worst recent events that impacted infrastructures and society, that at the time, pointed out the necessity of this studies to be born, then a brief explanation is given upon this topic. on the fourth chapter is approached the main topic of this paper, that is the convergence between sustainability and risk and resilience concepts in civil infrastructures, where is presented the current situation of research within these two areas, through a brief state of the art focused towards civil infrastructures, emphasizing on the necessity to fill this research gap and stating some key points that the creation of a methodology that addresses sustainability, risk and resilience for civil infrastructures must have. the last chapter are the conclusions, where the authors summarize the key points obtained from the research. 2. sustainable construction 2.1. history, definition, and application climate change has been increasing in recent years, being the result of the existence and the multiple activities conceived by humanity. among these activities, specifically, there is one of those that produce the greatest impact, which is the construction industry. this industry currently uses about 40% of the fossil fuels, 30% of the raw materials and 25% of the water consumed annually in the world. additionally, it consumes about 30% of electricity and generates 30% of greenhouse gas emissions [4], [9]. however, this industry has the potential to generate large reductions with small changes in the way buildings are designed and used, which could achieve savings of up to 30% in energy consumption, a 35% reduction in co2 emissions, and a reduction in water consumption of up to 50% [3, 4, 9]. this brought the attention to the international community to focus on the construction sector and its impacts began to emerge in the early 1970’s when the first united nations conference on environment and development was held in stockholm, giving rise to the united nations environment program, with the aim to promote the appropriate use and sustainable development [3, 4]. then, in 1988, in the report of the world commission on environment and development named "our common future", defined sustainability as "the attempt to meet the needs of the present without compromising the ability of future generations to meet their own needs". thus, it was only by the year 1992 that international principles for sustainable development were established, during the united nations conference on environment and development in rio de janeiro, where agenda 21 was elaborated, a document that systematized a plan of oriented actions creating the minimum necessary conditions for new constructions, both for developed and undeveloped countries [3, 10]. this rio conference has been continued with the versions of rio+10 in 2002 in johannesburg and rio+20 in 2012 in rio de janeiro, to reaffirm the rio-92 goals and to include clean energy and corporate responsibility in the debate and also to focus on the green economy, being capable of generating jobs with low impact on the environment and efficient use of natural resources [3, 11]. after these world conferences, the need to develop methods and tools to study the sustainable performance of buildings emerged, because the countries that were implementing projects with better environmental performance had no ways to verify their improvements, obtaining situations where green buildings consumed more energy than conventional ones. accordingly, it was necessary to develop methodologies to standardize building sustainability in a global way, allowing the analysis and comparison of various solutions to further improve the environmental performance of buildings [3]. 2.2. methodologies to assess sustainability currently there are tools that are not legally required, but they raise awareness and help to promote sustainability in construction entities. countries such as united states, canada, france, japan, among others have already implemented tools for assessing the sustainability of buildings, as a support for project design and at the same time assessing their postoccupancy. 618 vol. 33/2022 resilience and sustainability in civil infrastructures some of the first certifications of the sustainability level of buildings began to emerge through the use of methods such as breeam in the united kingdom, hqe certification (haute qualité enviromentale) in france, leed certification in the united states and gbtool (green building assessment tool) in some of the eu member states and adapted towards each state’s necessities [3, 4][3]. some of these methods are briefly explained down below. 2.2.1. leadership in energy and environmental design (leed) [3] developed by the u.s. green building council in 1994, in the united states of america. in this method, the environmental performance of the building is evaluated holistically, throughout its life cycle, i.e., in the design, construction, operation and maintenance phases. this tool applies to various types of buildings, such as residential, commercial, and school buildings, among others. for a building to be evaluated it must meet a minimum required criterion, a pre-selection with a series of requirements. after this verification, the building becomes eligible, and the stage of analysis and evaluation of its performance begins. this is a points-based system, awarding points for specific criteria in five different categories: sustainable sites; water efficiency; energy and atmosphere; materials and resources; and indoor environmental quality. these categories in total add up to 100 points, but there are two additional categories as a bonus, the innovation in design category, with 6 points, and the regional priority with 4 points. the points of the evaluation can only be assigned if the building complies with the requirements of the system; in the end, with the sum of these points, a classification can be assigned to the building, among four possible certifications, basic, silver, gold, and platinum. 2.2.2. building research establishment environmental assessment methodbreeam [3] the breeam method was developed in the early 1990s in the uk by researchers. the assessment with this tool considers the following sustainability categories: water, energy, materials, health and wellbeing, management, transportation, waste, contamination, and land use; each performance is verified by comparing them with pre-established benchmarks, thus obtaining the building’s assessment. breeam can be used for virtually any type of building, such as offices, industrial plants, residential buildings, and hospitals. thanks to the multiple existing versions of this tool, each one specifically developed to suit the building under assessment. in each of the categories mentioned above, requirements are defined, giving credits to the building, which are added as the building complies with them. on the other hand, in each category, specific weights are fixed, indicating their relevance, and depending on the type of building being evaluated. in this way, the set of credits and weights of the categories conform an environmental performance index for the building, which can obtain a value between zero and 100 for the environmental performance index (epi). according to the obtained epi, six levels of certification are attributed, "excellent", "great", "very good", "good", "approved", and "no classification", each one depending on the total points obtained in the environmental performance index. 2.2.3. sustainable building tool portugal sbtool pt [3, 4, 11] . this methodology follows four steps: (i) quantification of building performance at the level of each indicator; (ii) parameter normalization; (iii) parameter aggregation; (iv) sustainability score calculation and overall assessment. the overall rating of the building depends on the weighting of the different criteria, considering benchmarking practices reference practices that are set at the national level (portuguese level). this weighting is done by comparing the performance of the building with national reference practices: the best practice, which has a value of 1.0, and the conventional practice, with a value of 0.0. the value of this normalization is within a range between -0.2 and 1.2. subsequently, with the performance of each indicator and its corresponding weight, the performance of the category to which it belongs is calculated. finally, with the performances of each category and their respective weights, the results for each dimension are calculated, thus obtaining in the end an overall sustainability level for the building chosen for the case study. 3. risk and resilience analysis in critical infrastructures the growing number of catastrophic events and their implications, such as 9/11 in the united states, or the terrorist events in madrid on march 11, 2005, have prompted europe and other nations around the world to take steps to prevent these events from leading to high consequences that can be reduced or prevented. in june 2004, the european council called for a strategy to protect critical infrastructure, which had a response on october 20, 2004, in a statement in which the commission described the actions being taken to protect critical infrastructure and proposed additional measures to strengthen existing instruments and meet the mandates of the european council [12]. later, on november 17, 2005, the commission adopted a green paper on a european programme for critical infrastructure protection, stating that the main scope of the european programme for critical infrastructure protection (epcip) is the need to increase the capability to protect critical infrastructures (ci) in europe and to help reduce vulnerabilities related to ci [13]. 619 o. j. urbina, e. teixeira, h. sousa, j. matos acta polytechnica ctu proceedings the first step in protecting ci involves identifying and assessing the factors that may negatively influence its operations, defining a systematic, analytical approach to prioritizing resilience measures for ci. this analysis should include an assessment of the impacts of ci disruption by pre-established criteria. several approaches are used in oecd countries [14]. in terms of criteria, the european commission defines a minimum set for the assessment of critical infrastructure, including public impacts, economic impacts, environmental impacts, interdependence, political impacts, and psychological impacts. identifying weaknesses allows prioritizing where to focus resilience efforts in existing infrastructure systems: on points of failure that would have the most severe consequences. this prioritization can be a decisive variable in decision making, such as which infrastructure should be hardened or relocated, or which ci should receive priority restoration after a disaster to ensure rapid recovery [12, 14]. infrastructures must be designed and built to serve adequately over longer life terms without major deterioration or reaching collapse. with the current technology of materials, analysis, design, and construction, it should be possible to specify a bigger design life in new infrastructure projects. this leads towards a broader perspective which indicates that risk and resilience assessments are part of the general decision support to plan, to design and produce new infrastructures that are economically efficient, reliable, safe, secure, and sustainable [15]. 4. integrated approach for civil infrastructures recent literature has discussed the importance to go beyond the sustainability assessment of single buildings and to enlarge the assessment scale to communities and cities to meet all the different aspects of sustainability [9]. this chapter is divided in two main parts, the first part, where are presented some of the tools used to assess sustainability in infrastructures, and the second part, present diverse studies that integrate risk or resilience assessments with at least one dimension of sustainability’s assessment (environmental, social, or economic). 4.1. sustainability in infrastructures sustainability assessment tools for civil infrastructures are less frequent rather than the existing tools to evaluate buildings, however, some researchers have already progressed in this topic. following are showed some of their studies. • pardo-bosch et. al. [16], presented the multicriteria decision system mives. this decision model is divided into 4 steps. i) identification of the problem; ii) development of the decision tree, a diagram that organizes and structures the concepts that will be evaluated; iii) defining the relative weight of each of the aspects that are to be considered in the decision tree using analytical hierarchical process; and iv) establishment for indicators of value function that in each case reflects the appraisal of the decision-maker [16]. it prioritizes with technical accuracy public infrastructure projects that one administration must finance with only one budget in a developed country, helping to minimize the subjectivity in the entire decision-making process. • rosasco et. al. [8], show a study under a program named "regional strategic intervention program" (p.r.i.s), which objective is to guarantee the social protection of citizens that reside according to the italian law about the expropriation of private real estate for the construction of public work projects. within the framework of this program, the authors developed a mass appraisal estimation model that quantifies the indemnities values through a multi-parameter model of residential and commercial units within the area affected by the public work project. this multi-parameter estimation model uses a survey evaluated on the real estate units that are in the studied area; seven main features were selected, which are: 1. dimension (sqm); 2. age of building; 3. type of building; 4. maintenance state; 5. floor level; 6. lift (or not); 7. accessibility. this estimation model was applied in a new infrastructure project located near the city of genoa, in italy, the results show that the acceptance percentages of the indemnities estimated present a high degree of satisfaction, although the indemnities partially compensate all the inconvenience suffered. the authors conclude that these economic indemnities contribute to the achievement of that social sustainability of the infrastructures, and it can guarantee reasonable transfer alternatives for both residence and economic activities. • jones et. al [17], considered the holistic nature of water infrastructure development in terms of rural areas in developing countries, through an outcomebased assessment method using life-cycle analysis (lca). this method uses the lca framework to supply a holistic set of sustainability indicators, these indicators are grouped into three categories including metrics for technical (performance), environment, and economic (market-based) aspects. all the metrics are employed at community level to indicate the current state of the system. finally, this method should provide the socio-economic impacts generated. the authors applied the method on a generic example of arsenic water-treatment in bangladesh. 4.2. risk-resilience and sustainability in civil infrastructures the evaluation of resilience should not only consider technical but also environmental, organizational, social and economic dimensions [9]. despite this, there 620 vol. 33/2022 resilience and sustainability in civil infrastructures is limited number of previous studies that integrate them with critical infrastructure sustainability. few authors have worked onto developing a comprehensive approach that unites sustainability and riskresilience analysis as it could provide synergies and clearer results about the overall role and needs of civil infrastructures. • markert et. al. [15], presented a method for risk and sustainability analysis of complex hydrogen infrastructures, this model is based on a high level risk assessment, that are complemented with other decision support tools such as gis, lca and life cycle cost (lcc). providing a novel and comprehensive study as it enables the possibility to obtain spatial analysis with the gis system, for both, the risk assessment approach, and the sustainability approach, easing the identification of vulnerable elements and high-risk zones, in addition to the environmental and economic aspects. • bocchini et. al. [18], concluded that sustainability and resilience have a great deal of similarities and common features, as i.e. both integrate structural analysis with social and economic aspects, and both seek to enhance an infrastructure in terms of structural design, used materials, maintenance, management strategies, and impacts on society. therefore, the authors proposed an approach established on a risk assessment, using the concepts of probability of occurrence and risk, to address resilience and sustainability at the same time. despite the differences between sustainability and resilience targets, both converge on seeking to perform a service level to society during and after the occurrence of an extreme event and recover to the optimal functionality at great pace. for this, the proposed approach must be rigorous, quantitative, and unified, being able to assess numerous events and compare them. these events are weighted using eq. 1 to quantify their probability of occurrence. i = ! e∈er peie + ! e∈es peie (1) where i, is expected life-cycle impact of the infrastructure under analysis on the community (in monetary terms), and are the domains of events addressed by resilience and sustainability, respectively; is the probability of occurrence of the event e, and is the predicted impact on the community of the event e, this result is nondimensional. eq. 1 can be used at the individual scale of an infrastructure (e.g. a bridge), or at a global scale. • mejia et. al. [19], defined three performance measures for the long-term investment in energy and transport infrastructures, which are: cost, sustainability and resiliency. the authors defined sustainability as the period where the infrastructure system satisfies operational and environmental requirements, and resilience as the cost from a precontingency state of the infrastructure to a postcontingency due to an occurrence of an extreme event. in addition, the authors used four input factors to quantify the resilience, cost, and sustainability performance of the system: the system of interest; the expected demand; projected events, and the expected actions. this quantitative approach was demonstrated in a case study of an integrated energy and transportation system in the united states. • marinella giunta [20] proposed an integrated approach of resilience and sustainability to identify the most efficient alternative of road infrastructure rehabilitation after an extreme event based on life cycle costs. the author addressed that the main difficulty of this kind of approach is to do an evaluation of the sustainability and resilience using quantitative indicators. to overcome this issue, she articulated the approach in three main steps: 1. identify the rehabilitation alternatives considering technical, economical, and time aspects. 2. quantify the life cycle cost of each alternative. for each alternative it is quantified the life cycle assessment (lca) and the life cycle cost analysis (lcca). 3. resilience assessment of the infrastructure for each rehabilitation alternative, based on economic aspects. it is done by the estimation of the costs to restore functionality, and the time needed for the reconstruction. the result of this unified approach is a best solution of rehabilitation based on the lowest sum of the costs of sustainability and resilience. 4.3. conclusions this paper provided a broad analysis of the characteristics of different tools, methodologies and applications for the identification and evaluation of risks, resilience, and sustainability in civil infrastructures, as an opening towards developing an integrated and multidisciplinary methodology that allows to assess simultaneously to serve adequately over longer life terms without major contingencies. it was observed in the literature from the most relevant articles related to risk and resilience assessment, that there is a possibility to create a multidisciplinary methodology that includes sustainability as it directly involves the economical and societal consequences that any disruptive event might bring. in this way, the aspects associated with sustainability and resilience assessment are considered simultaneously and in a process of mutual interaction, opening the possibility to be used for planning and development of new cities, industries, and facilities as this comprehensive analysis is important for their continues optimization. the development of this kind of 621 o. j. urbina, e. teixeira, h. sousa, j. matos acta polytechnica ctu proceedings research seems to push designers and administrators in a similar direction, as it is a useful source of information for stakeholders at the local level, as the assessment methods and tools may comprehensively support sustainable choices that are also the basis for the sustainable pursuit of civil infrastructures. risk, resilience, and sustainability are complementary characteristics for civil infrastructure. while sustainability addresses the time-continuous impacts on the economy, society, and environment that the infrastructure certainly will distribute over its entire service life, resilience, and risk focus on the big impact that the service failure of the infrastructure can have in the case of extreme events. the combination of these approaches will provide a truly comprehensive assessment of the quality of the infrastructure. a possible big obstacle for the integrated assessment of both sustainability and resilience is the computation of truly quantitative metrics. the resilience research is more advanced in terms of quantitative analyses and indicators for civil infrastructures. instead, sustainability assessment systems have promoted a culture of qualitative assessment. finally, it is essential to continue investigating and improving the way of collecting and processing these methodologies. hence, future studies are needed to conduct to validate the usefulness and reliability of a comprehensive methodology described to evaluate the resilience and sustainability of civil infrastructures. additionally, it is vital to realize applications on case studies embracing different scenarios for each characteristic. acknowledgements this work was partly financed by feder funds through the competitivity factors operational programme compete and by national funds through fct (foundation for science and technology) within the scope of the project poci-01-0247-feder-039555). references [1] j. a. barbosa, l. 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[19] d. mejia-giraldo, j. villarreal-marimon, y. gu, et al. sustainability and resiliency measures for long-term investment planning in integrated energy and transportation infrastructures. journal of energy engineering 138(2):87-94, 2012. https: //doi.org/10.1061/(asce)ey.1943-7897.0000067. [20] m. giunta. sustainability and resilience in the rehabilitation of road infrastructures after an extreme event: an integrated approach. the baltic journal of road and bridge engineering 12(3):154-60, 2017. https://doi.org/10.3846/bjrbe.2017.18. 623 https://doi.org/10.1061/(asce)is.1943-555x.0000177 https://doi.org/10.1061/(asce)ey.1943-7897.0000067 https://doi.org/10.3846/bjrbe.2017.18 https://doi.org/10.14311/app.2022.33.0578 acta polytechnica ctu proceedings 33:578–584, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague load transfer mechanism of concrete screws florian stocker, oliver zeman∗, michael schwenn, konrad bergmeister university of natural resources and life sciences, department of civil engineering and natural hazards, institute for structural engineering, vienna, austria ∗ corresponding author: oliver.zeman@boku.ac.at abstract. even though the market and development for concrete screws has been increasingly rising in recent years, the load transfer mechanism of concrete screws has not yet been fully investigated. therefore, different tests of concrete screws made of galvanized and stainless steel were performed in concrete c20/25 and c50/60. the main aim is to measure the strain along the embedment depth. this will be achieved by using strain gauges that get placed in a centrically drilled borehole through the concrete screw. to get a comparison to the mechanism of the screws the same process will be executed in threaded rods used as a part of bonded anchors. due to the fact that the threaded cuts of concrete screws have geometrical similarities to bonded anchors, it was examined if the load transfer of both fasteners is related and may be compared. the results of the testing have shown that the load transfer mechanism of both fastener types is similar in low-strength concrete showing a concrete cone failure. in high strength concrete due to the mainly occurring steel failure the maximum measured strains at the maximum load step are not comparable. however, at lower load steps where the steel does not exceed the yield strength the results show a similar load transfer mechanism, too. keywords: anchorage to concrete, concrete screw, deformation behaviour. 1. introduction in fastening technologies there are three basic principles of load transfer mechanisms: friction-based, bond-based and mechanical interlock-based load transfer. the load transfer mechanism of postinstalled concrete screws is based on the principle of the mechanical interlock as shown in [1, 2]. concrete screws typically consist of five different segments: the head, the shaft, the run-out of thread section, the main load-bearing section and the thread cutting section. to provide the form locking function, the specially hardened thread cutting section is liable for cutting the threads for the following ones into the base material. the research presented in this paper is based on the question of how the load transfer behaviour behaves from the tip of the concrete screw to the upper edge of the concrete. this has already partially been examined by lechner et.al. [3] for concrete screws with a larger diameter used as retrofitting for shear force strengthening. therefore, a hole was drilled through the concrete screws along the embedment depth. in this manufactured hole, strain gauges got placed in specified intervals to measure the strain along the screw during a defined load situation. furthermore, to compare the load transfer mechanism of concrete screws with bonded anchors, the same process was done for threaded rods that get placed in the borehole with an injection mortar. 1.1. scope in the european assessment document ead 33023200-0601 "metal anchors for use in concrete" [4], concrete screws have different segments with specified tasks, see figure 1. the head (1) for transferring the load into the thread sections. the shaft (2) is for attaching components. it is dimensioned that it can be completely sunk into the borehole. the runout area where the thickness of the thread is running out to zero. the main load bearing section (4) is supposed to bear the main load into the base material and the thread cutting section (5) cuts the threads into the base material. so, every section has its specific task, however the detailed load transfer behaviour is not fully examined yet. it is assumed that the load transfer mechanism is similar to them from bonded anchors because both systems are transferring the loads into the basement almost over the entire embedment depth. the tests of the concrete screws presented in this contribution have been performed with an unconfined setup for an unrestricted concrete cone failure according to technical report tr 048 "details of tests for post-installed fasteners in concrete", [6]. this paper wants to provide a look into the segments of a concrete screw and their reaction with an applied load. 1.2. load transfer mechanism of concrete screws and bonded anchors the mechanical interlock principle is providing the load transfer between the fastener and the base material. this principle is used by different fastener systems, for example drop in anchors, headed anchors, undercut anchors, concrete screws and any type of 578 https://doi.org/10.14311/app.2022.33.0578 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 load transfer mechanism of concrete screws ���� ���� ���� ���� ���� � figure 1. sections of a concrete screw; modified according to [4]. figure 2. load transfer mechanism of concrete screws (left) and bonded anchors (right) based on the approach of [5]. cast-in anchorages. concrete screws in general are cutting the thread in the concrete through the hardened thread cutting section. figure 1 shows a main load-bearing section ((4) in figure 1) in which the main load is supposed to be transferred into the base material. furthermore, it is estimated that every single helix is contributing to the load transfer because all threads are fixed in the concrete. however, it is still not completely clear how much each thread is included in the load transfer mechanism. the question is which threads are contributing more to the load transfer into the base material. the strains of the two concrete screw types and the bonded anchor may differ from each other because of different material characteristics like e-modul, yield strength, tensile strength and fracture strain. the mechanical and the chemical interlocking principle are depicted in figure 2 schematically based on the approach given in [5]. the left picture in figure 2 shows the concrete screw interlocked in the base material loaded with a force and transferring the load through the threads. the picture on the right shows the bonded anchor and the theoretical load transfer through the bond stress. the failure modes are well described in the relevant literature [1, 7]. two of the failure modes for mechanical post-installed fasteners are concrete cone failure or combined failure. in this research there have been used single concrete screws with a nominal embedment depth of hnom = 85 mm and an effective embedment depth of hef = 68 mm. the calculation for the design situation of a single concrete screw follows the concrete capacity design (ccd) approach [8] where the ultimate strength of the anchor is mostly based on the square root of the concrete compressive strength and the embedment depth to the power of 1,5 [7, 9]. the load transfer mechanism of bonded anchors is working with chemical interlock. the different types of bond systems and how to work with them are well described in [1, 7]. the effective embedment depth of the bonded anchors has been hef = hnom = 85 mm in this research. the main bond models are the uniform bond model and the elastic bond model [7]. the uniform bond model is mainly dependent on the strength of the anchor (the anchor itself and the injection mortar), the embedment depth and the diameter [7]. it assumes a constant stress distribution along the embedment depth. if the bond stress surpasses the bond strength, then failure occurs. for high loads and strength design methods this model is more appropriate than the elastic bond model [7]. the bond stress τ can be derived by testing based on the following equation. nu = τ · π · d · hef (1) where • nu . . . ultimate strength of the anchor, • τ . . . bond stress, • d . . . diameter of the fastener, • hef . . . effective embedment depth. cook et. al. [7] showed that for low loads the elastic bond model is more appropriate than the uniform bond model to describe the anchor behaviour. a finite element analysis made by mcvay, cook and krishnamurthy [10] also showed that bonded anchors following the elastic bond model at low load levels and the uniform bond model at peak loads. following calculation can be made based on these results. nu = τ · π · d0 ! tanh (λ · hef ) λ " (2) 579 f. stocker, o. zeman, m. schwenn, k. bergmeister acta polytechnica ctu proceedings figure 3. test samples and position of the strain gauges. type concrete strength n test samples 20−a c20/25 5 galvanized steel concrete screw 20−b 3 stainless steel concrete screw 20−c 3 threaded rod 10.9 50−a c50/60 5 galvanized steel concrete screw 50−c 3 threaded rod 10.9 table 1. overview of the testing program. where • d0 . . . diameter of the borehole, • λ . . . stiffness characteristic of the bond anchor system. davis and cook showed with a finite element analysis of an elastic model that the behaviour of a bonded anchor with a low short term load has a stress distribution relationship like an hyperbolic tangent curve [4]. this hyperbolic tangent curve approaches a distribution of the bond stress like the uniform bond model as the load is increasing or the duration of the test increases at a specific load level [11]. the load transfer behaviour of the bonded anchors that was found in this research seems to reflect the behaviour of the elastic model also approaching a hyperbolic tangent curve. the load transfer principle and the calculation approaches are different from form-locking systems like concrete screws. however, through the threads along the screw it is assumed that bonded threaded rods have a similar bearing behaviour like concrete screws. 1.3. tested concretes the tests shown within this contribution have been performed in non-cracked concretes of strength classes of c20/25 and c50/60 from the same supplier, with a concrete compressive strength measured at cubes of fc,cube = 29, 4 n/mm2 for the c20/25 and fc,cube = 68, 2 n/mm2 for the c50/60 at time of testing. the maximum aggregate size (round aggregate) was defined by dg = 16 mm following the grading curve given in tr 048 [8] which is based on din 1045-2 [5]. the resistance to fragmentation of the aggregates is at a los angeles abrasion value of 28 and an impact value of 21 for the considered low strength concrete, according to [12]. 2. test program and description 2.1. general in standard pull-out tests the result is usually forcedisplacement relationship. but there is a lack of information about the behaviour of the screw segments along the whole anchoring depth and to what extent they are involved in the load transfer process. this paper aims to examine how the load transfer behaves and makes a comparison to the behaviour of bonded anchors. in order to make this behaviour visible, the strains along the concrete screws and the bonded anchors are measured by means of strain gauges. to develop the test rig, a centrically arranged hole with a diameter of 3,5 mm has been drilled through the concrete screw. the strain gauges (short called dms) are placed in the borehole through the testing samples as shown in figure 3. table 1 shows the test program and the test samples. therefore, the testing program consist of five centrically drilled concrete screws made from galvanized steel, three centrically drilled concrete screws 580 vol. 33/2022 load transfer mechanism of concrete screws ��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh� ���$ �����ordgvwhs ������ordgvwhs ������ordgvwhs ������ordgvwhs džm [‰] figure 4. type 20−a εm[‰] in 25 % loadsteps with concrete screws made of galvanized steel in c20/25; [n = 5]. made from stainless steel and three centrically drilled threaded rods with a steel strength of 10.9 in the concretes of c20/25 and c50/60. the measurements of the concrete screw are 9,4 mm for the core diameter, 12,5 mm for the total diameter including the thread and about 110 mm for the total length. the threaded rod has a nominal diameter of 10 mm. the exact value of the diameter for a m10 without the threads is 8,6 mm. the nominal embedment depth for all tests was set to 85 mm with hnom ≈ 9d with a diameter d = 9, 4 mm. for the hammer a medium drill bit size of the diameter was used. this means for the concrete screws dcut = 10, 25 mm − 10, 35 mm and for the bonded anchors dcut = 12, 25 mm − 13, 25 mm . the used cement-based injection mortar has a nominal bond strength of 18 n/mm2, the bonded anchors have been installed following the manufacturerťs product installation instruction. the strength of the screws has been tested on samples out of the middle part of the cross section. the measured tensile strength of the galvanized concrete screws was 960 n/mm2. therefore, to provide aăgood comparison between concrete screws and bonded anchors, threaded rods of strength class 10.9 with a tensile strength of 1000 n/mm2 have been chosen for the bond test sample. the screw made of stainless steel was chosen out of the same product line as the galvanized one, tests have shown that the tensile strength of the inner part is 780 n/mm2. 3. evaluation of the test results 3.1. test results in this section the results of the performed tests are provided. in the evaluation of the results, the measured strains of the 5 strain gauges are depicted. these strain curves are arranged in loadsteps of 25 % in the order of 25 %, 50 %, 75 % 100 %. for the type 20−a and 20−b a descending strain curve is depicted in figure 5 and figure 7. these are showing the strain at a load level of 75 % after the maximum load of 100 % was reached. the analysis is based on mean ��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh� ���$ �����ordgvwhs ������ordgvwhs�diwhu�pd[��ordg džm [‰] figure 5. type 20−a εm[‰] of 100 % and 75 % loadsteps after max load with concrete screws made of galvanized steel in c20/25; [n = 5]. values of the measured data, not considering the scatter. for further results, see [13]. the scattering of the strain values is wide. though, with an increasing distance from the tip to the head of the test samples the data shows more consistency as a result of lower scatter. 3.2. results of the concrete screw made of galvanized steel in concrete c20/25 in this section the results of the tests of galvanized concrete screws in concrete c20/25 are provided (type 20−a in table 1). figure 5 is showing the strain ε [‰]] along the embedment depth of the screw. on the left side, depicted are the strains in 25 % steps beginning with 25 % up to 100 % of the maximum load. the graph shown on the right is depicting the 100 % step and the 75 % step after the peak load of the fastener. the x−axis is providing the position of the strain gauges. for the positioning, see figure 3. the maximum mean load (n = 5) is fm = 46, 50 kn with a cv = 9,08 %. all fasteners in this test series failed with concrete cone failure. each thread of the concrete screw seems to be included in the load transfer with a specific amount. most of the load transfer appears to take place at positions 3 and 4, therefore, these measurements are consistent with the assumptions in figure 1 that the main load is transferred into the base material at the "main load bearing section". beginning with position 2, it can be stated that there is a linear increase in the strain up to the strain gauge at position 5. 3.3. results of the concrete screw made of stainless steel in concrete c20/25 the strain of the concrete screw (type 20-b) has higher values than that from the galvanized concrete screws (see the following graph shown in figure 6). at the same load level, the stainless steel screw stretches nearly double the size compared to the gal581 f. stocker, o. zeman, m. schwenn, k. bergmeister acta polytechnica ctu proceedings ��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh� ���% �����ordgvwhs ������ordgvwhs ������ordgvwhs ������ordgvwhs i̇p >å@ figure 6. type 20−b εm[‰] in 25 % loadsteps with concrete screws made of galvanized steel in c20/25; [n = 3]. ��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh� ���% �����ordgvwhs ������ordgvwhs�diwhu�pd[��ordg i̇p >å@ figure 7. type 20−b εm[‰] of 100 % and 75 % loadsteps after max load with concrete screws made of galvanized steel in c20/25; [n = 3]. vanized version. the maximum mean load at these tests is f = 45, 47 kn with a cv = 9,83 %. based on the characteristics of both steels, the higher strain of the stainless steel screw is a result of the higher fracture strain up to εb = 40 %. the failure mode of this test series was concrete cone failure too. furthermore, the cutting tip failed in the middle of the thread cutting section in 2 out of 3 tests. the graph is showing a test sample n = 3 and after analyzing the data, the peak at position 3 is due to the result of one experiment at a loadstep of 100 %. the other tests show an almost linear behaviour (dotted line in figure 7). with a higher number of tests, the strain curve could be harmonized (like the curves at type 20-a). 3.4. results of the concrete screw made of galvanized steel in concrete c50/60 the results of the concrete screw in a c50/60 concrete (type 50-a) show that up to the 75 % loadstep the distribution of the measured strains have a similarity to the type a tests in low strength concrete c20/25. the strain values are slightly higher, this may be caused due to a higher stiffness of the con��� ��� ��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh� ���$ �����ordgvwhs ������ordgvwhs ������ordgvwhs ������ordgvwhs džm [‰] figure 8. type 50−a εm[‰] in 25 % loadsteps, galvanized concrete screw in c50/60; [n = 5]. ��� ��� ��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh� ���$ �����ordgvwhs �����ordgvwhs�diwhu�pd[� ordg i̇p >å@ figure 9. type 50−a εm[‰] of 100 % and 75 % loadsteps after max load with concrete screws made of galvanized steel in c50/60; [n = 5]. crete. after the 75 % loadstep, 3 out of 5 tests failed due to steel failure of the concrete screw. the tensile force exceeded both, the yield strength and the tensile strength, which is the reason for the peak at position 4 in figure 8. the screws with steel failure broke at position 4 with a maximum load of 58,33 kn with a cv = 5,26 %. the 75 % loadstep after the maximum load in the graph shown on the right are the strains of the screws with concrete cone failure (n = 2), which show also a similar behaviour as the tests in low strength concrete. 3.5. comparison to the bond model in this section the results of concrete screws made of galvanized steel and the bonded anchors with threaded rods steel class 10.9 in a concrete c20/25 will be compared. the values of the measured strains of the bonded anchors (type 20-c) are higher than the strains at the galvanized concrete screw (type 20-a), see figure 10. but it turns out that the strain development over the embedment depth can be assumed as quite similar, as shown in figure 10. figure 10 is showing the strain curves of the type 20a and 20-c tests compared at each 25 % loadstep. up to thew 75 % loadstep the strain curves showing a similar behaviour (except the specific strain values 582 vol. 33/2022 load transfer mechanism of concrete screws ��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh����$�dqg����& ����ordgvwhs i̇p >å@ ����ordgvwhs�w\sh����$ ����ordgvwhs�w\sh����& --��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh����$�dqg����& ����ordgvwhs džm [‰] ����ordgvwhs�w\sh����$ ����ordgvwhs�w\sh����& -��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh����$�dqg����& ����ordgvwhs i̇p >å@ ����ordgvwhs�w\sh����$ ����ordgvwhs�w\sh����& --��� ��� ��� ��� ��� ��� ��� ��� � � � � � � 7\sh����$�dqg����& �����ordgvwhs i̇p >å@ --�����ordgvwhs�w\sh����$ �����ordgvwhs�w\sh����& figure 10. comparison of εm[‰] each 25 % of galvanized concrete screws (denominated as type 20-a) and bonded anchors (type 20-c). at positions 4 and 5). at the maximum load (the 100 % loadstep), shortly before the fasteners failed the strain curves of the 20-a and 20-c types differ from each other. but there are two different failure modes developing. the bonded anchors (type 20-c) failed by steel failure with a maximum load of 49,43 kn (cv = 11,13 %) where the maximum load of the concrete screws (type 20-a) was 46,50 kn (cv = 9,08 %) by concrete cone failure. as discussed previously, the elastic bond model is a good approach for low load levels before the peak [7]. this is also confirmed in figure 10. these strain curves increase evenly like in the elastic bond model. regarding the similarity of the steel strength and the yield strength of both anchor systems (960 n/mm2 to approximately 1000 n/mm2), the difference in failure modes can be a reason for the higher strain values in the bonded anchor. however, the graphs in figure 10 compared may represent an indication that the mechanical behaviour and the load transfer behaviour of concrete screws can be assumingly similar to bonded anchors. lastly, the shape of the curves shown in test series type 20-a and 20-c is a confirmation of the elastic bond model discussed before. the comparison between the results of the bonded anchor (type 50-c) and the concrete screw (type 50-a) in a c50/60 will be left out. the strain curve of the bonded anchor (type 50-c) is similar to the bonded anchor (type 20-c), and the strain values are slightly higher. however, the different failure modes of the concrete screw (type 50-a) in a c50/60 does not allow a practical comparison. 4. conclusion in this study, tests with concrete screws and bonded anchors were performed in uncracked concrete c20/25 and c50/60. all tests had a nominal embedment depth of hef = 85 mm. regarding the results of the concrete screws, the main load bearing section is carrying most of the load into the base material. in this test series the main load bearing section of concrete screws made of galvanized steel transferred 90 − 95 % of the load into the concrete, depending on the concrete strength. the higher the strength of the concrete is, the more strains got measured in the main load bearing section. this is confirmed by the measured strains at dms positions 3 and 4. this also reflects the failure mode of the type 50-a tests where steel failure only occurred at dms position 4. the thread cutting section seems to be included in the load transfer mechanism. the values of the measured strains at dms positions 1 and 2 are low compared to the dms positions 3 and 4, but nevertheless this section is also able to bear a part of the load into the base material. (5 − 10 %) the amount seems to be influenced by the screw material. the stainless-steel screw has higher strains at the tip as the galvanized concrete screw. in this test series the concrete screw made of stainless steel was with up to 15 % included in the load transfer mechanism. 583 f. stocker, o. zeman, m. schwenn, k. bergmeister acta polytechnica ctu proceedings the results of the data indicate that the load transfer behaviour of concrete screws in concrete c20/25 are similar to bonded anchors. the strain curves of the tested bonded anchors are comparable to the elastic bond model [7]. therefore, as seen at the strain curves, concrete screws can be assumed to react like a bonded anchor (a threaded rod) with the elastic bond model. the behaviour of concrete screws investigated in this test series is only valid for the parameters used to perform these tests. therefore, the failure modes should be considered. the compared threaded rods have failed with steel failure and the concrete screws with concrete cone failure. however, before the steel failure occurs, the measured strains or at least the form of the curves can assumingly be compliant. the results themselves appear to be satisfactory, but additional tests should be performed to (1) enlarge the number of samples and (2) observe additional load transfer mechanisms of different concrete screw anchors made from different materials. references [1] r. eligehausen r., r. mallée, j. silva. anchorage in concrete construction, berlin: ernst & sohn, p. 370, 2006. [2] j. küenzlein. tragverhalten von schraubdübeln unter statischer zugbeanspruchung dissertation (stuttgart: selbstverlag des autors), p. 223, 2005. [3] j. lechner, n. fleischhacker, c. waltl, et al. zum verbundverhalten von betonschraubdübeln mit gro)em durchmesser. betonund stahlbetonbau 112(9):589-600, 2017. https://doi.org/10.1002/best.201700043. [4] eota. ead 330232-00-0601 mechanical fasteners for use in concrete (brussels: european organization for technical assessment), p. 57, 2016. [5] din 1045-2. tragwerke aus beton, stahlbeton und spannbeton teil 2: beton festlegung, eigenschaften, herstellung und konformität anwendungsregeln zu din en 206-1, berlin: deutsches institut für normung) p. 48, 2008. [6] eota. technical report tr 048: details of tests for post-installed fasteners in concrete (brussels: european organization for technical assessment), p. 26, 2016. [7] r. a. cook, j. kunz, w. fuchs, et al. behavior and design of single adhesive anchors under tensile load in uncracked concrete. structural journal 95(1):9-26, 1998. [8] r. eligehausen, j. küenzlein. zu: tragverhalten von befestigungen mit schraubdübeln. betonund stahlbetonbau 97(4):a16-a, 2002. https://doi.org/10.1002/best.200200980. [9] austrian standards institute. eurocode 2, en 1992-4: design of concrete structures, part 4: design of fastenings for use in concrete (vienna: austrian standards institute), p. 129, 2019. [10] r. a. cook. behavior of chemically bonded anchors. journal of structural engineering 119(9):2744-62, 1993. https://doi.org/10.1061/ (asce)0733-9445(1993)119:9(2744). [11] t. m. davis, r. a. cook. sustained load performance of adhesive anchor systems in concrete. aci structural journal 114(4), 2017. https://doi.org/10.14359/51689723. [12] austrian standards institute. önorm en 1097-2: 2010-03: tests for mechanical and physical properties of aggregates (vienna: austrian standards institute), p. 33, 2010. [13] f. stocker. untersuchungen zum lastabtragungsmechanismus von betonschrauben diplomarbeit, wien: universität für bodenkultur, p. 161, 2019. 584 https://doi.org/10.1002/best.201700043 https://doi.org/10.1002/best.200200980 https://doi.org/10.1061/(asce)0733-9445(1993)119:9(2744) https://doi.org/10.14359/51689723 doi:10.14311/app.2021.30.0104 acta polytechnica ctu proceedings 30:104–108, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app fire protection of steel elements using lightweight hybrid cement mortar stanislav šulca, ∗, vít šmilauera, jakub šejnab, františek waldb a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b czech technical university in prague, faculty of civil engineering, department of steel and timber structures, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: stanislav.sulc@fsv.cvut.cz abstract. material properties of steel structures are significantly reduced at high temperatures, so a fire protection has strong positive impact on the fire resistance of the structure. fire resistance of steel elements can be increased using a layer of cement-based materials as a fire protection. most of commonly used cement-based materials do not withstand high temperatures without noticeable reduction of mechanical properties. hybrid cement showed some interesting properties in the way of resistance to high temperatures and adhesion to steel surfaces, thus its behavior during fire exposure should be investigated. one experimental analysis with numerical simulation is presented in this article. it examines thermal material properties of lightweight hybrid cement mortar with expanded perlite from a simple experiment with a lab gas burner. keywords: fire protection, fire resistance, hybrid cement. 1. introduction mechanical behaviour of steel structures is very sensitive to high temperatures [1, 2] and it is essential to protect steel structures from fire exposure. there are various ways of such protection based on creating physical barrier with low thermal conductivity, such as fire-protective paintings, gypsum plasterboard lining or protection using cement-based materials. this article aims at a relatively new material called "hybrid cement" (h-cement) [3], produced by company "považská cementáreň, a. s." in slovakia. it has ambitions to be used as a fire-protective material, because the preliminary testing of this material by daxner [4] showed good characteristics at high temperatures compared to commonly used cementbased materials which do not withstand high temperatures without cracking and spalling [5, 6]. another one of its advantages is its suitability for steel structures, because it chemically binds to the steel surface, which makes noticeably stronger connection compared to other cement-based materials, but the adhesion itself has not been measured yet. there were several initial tests made to inspect the material properties. one of these tests is presented in this article. the overall target is to determine advantages of this material, explore and test its possible usages in the structures. 2. lab burner test the thermal characteristics of the material during fire exposure can be calculated by analysis of experimental data. as classic experiments in big furnaces are very expensive, the need of more experimental data lead to a small experiment with a lab gas burner. 2.1. experiment description the subject of interest is a h-cement mortar with expanded perlite. the target of this experiment was to determine its thermal conductivity. the composition of this light-weight h-cement mortar is in table 1. the amount of expanded perlite was composed of 43 % of fraction 0-1 mm and 57 % of fraction 12 mm. bulk density of the mixture in fresh state was 437 kg.m−3, while in the state of the experiment it was 335 kg.m−3. the bulk density is highly dependent on the level of intensity and duration of mixing, which mills the granulas of expanded perlite, see figure 1. it may not be the most suitable material for this purpose due to very small mechanical resistance. h-cement 72.1 g expanded perlite 9.17 g plastifier "stachement 508" 1,11 g water 55,6 g table 1. composition of h-cement mortar with expanded perlite. a steel element of 12 mm thickness was covered with h-cement mortar with expanded perlite from the bottom side. the other sides were covered with mineral wool. see the geometry scheme in figure 6 and the photos in figures 2-5. the hot air coming from the bottom of the element was being dissipated using thin steel plate aligned with the bottom side of the element. 104 http://dx.doi.org/10.14311/app.2021.30.0104 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 fire protection of steel elements using hybrid cement figure 1. comparison of regular (left) and careful (right) mixing of the mortar /the weight of both elements is the same/. figure 2. experiment setup. figure 3. experiment setup. the properties of used materials are summarized in tab. 2. before the experiment was realized, the mortar was being dried in 40 ◦c for 9 days, during which the mortar lost 30 % of its weight. in the 9th day the weight losses were about 0.8 % per day. the lab gas burner was placed under the center of the exposed surface. the temperatures were measured with three thermocouples mtc 10 type k. the first one measured the temperature of the 12 mm thick steel element. the second one was placed on the center of the exposed surface. the third one was figure 4. experiment setup. figure 5. experiment setup. h-cement mortar bulk density 335 kg.m−3 capacity 1150 j.kg−1.k−1 conductivity unknown steel bulk density 7850 kg.m−3 capacity 500 j.kg−1.k−1 conductivity 46 w.m−1.k−1 mineral wool bulk density 100 kg.m−3 capacity 800 j.kg−1.k−1 conductivity 0.041 w.m−1.k−1 table 2. simulation material parameters. also on the exposed surface, but positioned close to the corner of the surface, see the photo in figure 3 and a position scheme in figure 7. all the three temperature curves are depicted in figure 9. we hoped in reaching higher temperatures, but the thicknesses of the mortar and the steel element were too big and the experiment was stopped after 45 minutes, thus the temperature of the steel element reached only 167 ◦c. but the layer of h-cement was in a state of high temperature, see the simulation results in figure 11. immediately after the 45 minute fire load the mor105 s. šulc, v. šmilauer, j. šejna, f. wald acta polytechnica ctu proceedings figure 6. geometry of the experiment. figure 7. position of the thermocouples on the exposed bottom surface. 3×3 fem elements are indicated. tar was put into water to test its mechanical resistance for this kind of immediate change of temperature, while the steel element remained just above water. the state of the mortar is shown in figure 8 and shows no visible damage. 2.2. simulation the simulation was realized with oofem software [7]. the domain was composed of quadratic brick elements. the domain of steel member was composed of 3 × 3 × 1 fem elements. the domain of h-cement was composed of 3 × 3 × 4 elements. the thickness of mineral wool was divided into 3 elements. material properties follow values from table 2. the temperatures ast center and ast side are considered to be the "adiabatic surface temperatures" figure 8. h-cement mortar after the experiment and after fast cooling in water. and thus come into both convective and radiative boundary conditions on the exposed surface. the temperature ast center is applied on the surface of the brick element in the center of the domain, while the rest of the exposed surfaces are loaded with the ast side temperature. the element division is noticeable in figure 7. the initial temperature of the whole domain was 28.6 ◦c. the outer temperature in boundary conditions of mineral wool surfaces was considered to be 30 ◦c. the simulation was calculated with all timesteps being 1 second. the emissivity of exposed surfaces was 0.8 and the heat transfer coefficient was 20.0 w.m−2.k−1. the output of this simulation is a curve representing temperature of the steel element for given conductivity value of h-cement. three curves for different values of conductivity are depicted in figure 10 to show the impact of this parameter on the results. 106 vol. 30/2021 fire protection of steel elements using hybrid cement figure 9. adiabatic surface temperatures from the thermocouples and temperature of the steel element. figure 10. temperature comparison for three values of h-cement conductivity. figure 11. temperature distribution at time 2700 s. according to the simulations, the thermal conductivity of this h-cement mortar is 0.09715 w.m−1.k−1. these results slightly do not correspond with experimental data until 100 ◦c temperature is reached, which is caused by water content and its evaporation. after that moment the results match the experimental data very well and also seem to have the same tendency for higher temperatures, which were not reached in this experiment. the temperature field in a section through the center of the domain at time 2700 s is presented in figh-cement mortar bulk density 335 kg.m−3 capacity 1150 j.kg−1.k−1 conductivity 0.09715 w.m−1.k−1 table 3. h-cement material parameters with conductivity yielding from the simulations. ure 11. it shows that despite the low resulting temperature of the steel element, the temperature of the 107 s. šulc, v. šmilauer, j. šejna, f. wald acta polytechnica ctu proceedings mortar has reached quite high values. however, it would need another experiment to show its behavior on its limit at much higher temperatures, but it should be completely different experiment setup with another fuel. 3. conclusions h-cement is a material with promising characteristics and resistance to high temperatures during fire. this article determined its thermal properties based on a simple experiment with a lab gas burner and it also showed its behavior during fire load with temperatures comparable with the standard temperature curve iso 834. in principle, once we know the thermal material properties, simulations can replace the experiments in the process of design of fire protection. but more experiments with higher temperatures must be done to have data about the material behavior at full possible temperature range. the material will be further tested and analysed to find its strongest sides and its most suitable applications. acknowledgements this research was funded by czech science foundation, grant 19-22435s "performance of structures with timber fire protection – multi-physics modelling" and also by czech technical university in prague, grant number sgs20/038/ohk1/1t/11. references [1] g. li, p. wang. properties of steel at elevated temperatures. advanced analysis and design for fire safety of steel structures 2013. doi:https://doi.org/10.1007/978-3-642-34393-3_3. [2] z. nie, y. li, y. wang. mechanical properties of steels for cold-formed steel structures at elevated temperatures. advances in civil engineering 2020. [3] p. martauz, i. janotka, j. strigáč, m. bačuvčík. fundamental properties of industrial hybrid cement: utilization in ready-mixed concretes and shrinkagereducing applications. materiales de construccion, 2016. [4] j. daxner. use of hybrid cement in fireresistant plasters. 88 khi/2016, 2016. [5] i. hager. behaviour of cement concrete at high temperature. bulletin of the polish academy of sciences, technical sciences 61, 2013. [6] j.-c. mindeguia, p. pimienta, h. carré, c. l. borderie. experimental analysis of concrete spalling due to fire exposure. european journal of environmental and civil engineering 2013. doi:10.1080/19648189.2013.786245. [7] oofem https://oofem.org, 2020. 108 https://doi.org/https://doi.org/10.1007/978-3-642-34393-3_3 https://doi.org/10.1080/19648189.2013.786245 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.32.0037 acta polytechnica ctu proceedings 32:37–41, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the importance of a combined 3d model of ground and geotechnical structure ivan vaníček∗, daniel jirásko czech technical university in prague, faculty of civil engineering, thákurova 2077/7, 160 00 prague 6, czech republic ∗ corresponding author: ivan.vanicek@fsv.cvut.cz abstract. the rapid progress of digitalisation in civil engineering is also happening in the branch of geotechnical engineering. bim is a part of this process. the mutual connection of a 3d model of ground with a 3d model of geotechnical structure can help to give a clearer view of the interaction of this structure and ground. at the same time it also makes possible better control for all participants in the construction process, particularly if all conditions of this mutual interaction are fulfilled. there is considered whether the main aim of the construction process, namely that the care devoted to the individual phases of the structure design and execution actually correspond to the risk with which this structure and ground are indeed connected. within this context a closer specification of the eurocode 7 of the second generation will be identified , especially as it relates to different models. keywords: 3d ground model, 3d model of geotechnical structure, bim, phases of geotechnical investigation, principles of sustainable construction. 1. introduction practical problems solved in geotechnical engineering have usually a 2d character. a typical example could be a flood barrier along a river or a long retaining wall. the design and execution of both of these geotechnical structures is based on the characteristic values of geotechnical parameters of the soil with which these structures interact. the problem is that on a relatively long section there may be a place where the geotechnical parameters used in the design are lower than expected and in this local place a failure may occur. in the case of a flood barrier (a dike, for example), a section 10 m long that breaks will in practice devalue the principle of the entire flood barrier, which can be as much as 10 km long. of the various types of retaining walls, the gabion wall has the lowest stiffness in the third dimension, especially in comparison with conventional concrete. therefore, it is easier to cause a local failure at the weakest point. when a 2d ground model is used, at least in the form of geological and geotechnical models of the subsoil, it is easier to identify these potentially dangerous places and thus prevent a loss of functionality. for earth structures of transport infrastructures, where it is a line construction, but where the third dimension is even more important than for a flood barrier (dike) because the width (third dimension) for higher fills or deeper cuts is more important, even though in reality the fault retains its local character, a typical example being the landslide on the motorway d8. practically the same applies to linear underground structures, as in the case of tunnels. from the point of view of the geotechnical construction itself, there is an interesting transition for earth and rock fill dams, for which in the ratio of the length of the dam to its height, it becomes more appropriate to use a 3d computational model. the reason is a more credible assessment of the potential danger of tensile cracking in the sealing of these dams. for foundation structures, given the current more significant use of underground spaces (as for underground floors), the solution of a foundation structure in 3d is almost a necessity. the current requirements for digitalisation in the construction industry, e.g. in the form of the use of bim (building information modelling), are thus met for geotechnical structures. the basic requirement is that the entire structure, i.e. not only the superstructure but also the subsoil and the foundation structures mediating the mutual interaction, be part of one common 3d model. the interconnection thus brings to the fore the possibility of comparing the quality (informative value) of the subsoil model (whether geological or geotechnical) with the complexity of the superstructure ,which can again be geotechnical, as already mentioned in earth structures or underground structures such as tunnels, galleries etc. at the same time, it meets the requirements of čsn en 1997 design of geotechnical structures and the requirements for subsoil (ground) models obtained during different phases of geotechnical investigation. 2. bim one of the most widely used definitions of the bim process is given by the american national committee for the bim standard, which states that bim is a digital description of the physical and operational characteristics of a facility. bim is a shared source of knowledge and information about this device, creating 37 https://doi.org/10.14311/app.2022.32.0037 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en ivan vaníček, daniel jirásko acta polytechnica ctu proceedings a reliable basis for decision-making during its life cycle, defined from the first concept to demolition [1]. the emphasis is on two basic factors: • it is a reliable basis for the decision-making process; • it is a long-term process – covering the entire life of a building – from the concept itself to a final demolition. next, the main attention will be focused on the first factor, which in principle leads not only to a safe but also an effective construction design. all basic participants in the construction process can be part of this decision-making process, starting from the investor, and then on to the implementer of the geotechnical investigation, the designer, the supplier/contractor, and finally to the approval body, which comments on the construction. however, the second factor is no less important, as it mainly allows for time sequence and the continuity of individual steps, thus ensuring gradual development and specification. however, we must not forget the implementation of the principles of sustainable construction, which are currently strongly promoted [2, 3]. this means the possibility of assessing financial efficiency not only in the construction phase, but also in the demolition phase (i.e. through the entire life of the building). therefore, lower demands on energy (plus co2 footprints), lower land consumption as well lower demands on natural materials (by replacing them with non-standard materials) should be evaluated when total price is being determined. the question of a reliable basis for a decision-making process and for the mutual cooperation of the main actors is present from the very beginning. in the first place, there is the investor’s idea of the structure and the initial information about the ground on/in which this structure is to be situated. today, even investors in supermarkets or warehouses understand that one standard, unified project of such a building cannot be put in place, at least without a minimum knowledge of the nature of the subsoil. in the same way, it is not possible to model the upper structure in 3d without relation to the ground/subsoil, respecting its replacement by various springs, as this is contrary to the basic principles of structural design, now clearly expressed in the system of eurocodes for structural design. finally, this is shown by the recommended name of eurocode 0: eurocode – basis of structural and geotechnical design, defining the principles for the design of all building and civil engineering structures, including geotechnical – draft pren1990, september 2020, which will replace en 1990: 2002 [4]. therefore, the importance of interconnected 3d models of the subsoil and the upper structure (whether it is a building or an engineering structure, including a geotechnical structure), is at the centre of current attention. 3. models time sequence the current valid version of eurocode 7 – čsn en 1997 in part 1997-2 [5] in annex b1 presents the following scheme for the stages of geotechnical/ground investigation of the foundation soil in the geotechnical design, the execution of work and the use of the geotechnical structure – see figure 1. it is clear from this scheme that two basic parts are important within the geotechnical investigatory process: • providing a geological model (ground -subsoil, borrow pit), the main task of which is to divide the ground into different lithological layers with approximately identical properties, including different discontinuities, and also including information on the groundwater level. • ensuring the implementation of field or laboratory tests in the field of soil and rock mechanics for individual lithological layers. the geotechnical/ground investigation report (gir) is the main output of this process of geotechnical investigation. note. en 1997 of the second generation uses not only the term geological model, but for this geological model supplemented by the results of field and laboratory tests, also uses the simple term ground model. although the english word ground has many possible meanings in czech, it can probably be expected that it will be translated persistently as a model of the subsoil. although for underground structures (tunnels, metro) it will probably be a more appropriate translation to say the rock environment model, and as well as for the earth structures to say the soil environment model. within the design of the geotechnical structure, two models come to the fore: • the geotechnical model – developed by the designer o the geotechnical structure after the evaluation of geotechnical parameters and coefficients. • the calculation model, when the analytical calculation model and the numerical calculation model come to the fore. after the elaboration of the geotechnical structure project, the selection of the supplier and the actual implementation of all works will follow. note. in en 1997 of the second generation the first model is labelled as "geotechnical design model", which for each lithological layer specifies representative/characteristic values of the geotechnical parameters, and which are subsequently used in the calculation model. from the point of view of time sequence, ground/geotechnical investigation becomes very important in respect to its phasing. most often, and this 38 vol. 32/2022 the importance of a combined 3d model figure 1. stages of ground investigation in geotechnical design, execution of works and exploitation of the structure. is also in the second generation en 1997, for significant structures connected with high risk a distinction is made between: • desk study – a search of all available existing information from the point of view of the requirements of the geological and geotechnical model this phase is connected with the investment plan. in principle, it is based on all available map data – from geological, engineering-geological, hydrogeological to map data of a geo-environmental nature. also it makes use of the database of survey points carried out so far in the area of interest. for significant public civil engineering structures, these documents can be used to evaluate various variants of location, with subsequent use for the eia process. therefore it is important, and this is supported by the cgts cice, that the data obtained so far can be used in such a form that would allow their quick placement in the 3d model of the subsoil for the first evaluation. • preliminary gi – the preliminary geotechnical/ground investigation – the investigation already uses survey methods, although only to a limited extent, which serves primarily to confirm the expectations obtained in the desk study and from the point of view of designing a geotechnical structure, it serves as a preliminary design for documentation for the zoning decision / project plan. • design (detailed) gi – detailed investigation, investigation for the geotechnical structure design – this is the most important phase, as the results and conclusions are directly usable in the design of a geotechnical structure for the documentation phase regarding a building permit. 39 ivan vaníček, daniel jirásko acta polytechnica ctu proceedings • supplementary gi – additional investigation is needed if during the execution phase larger deviations from the expected ones are found, which may affect both the actual technology of execution and any corrective calculation from the point of view of limit states. • confirmatory gi – a confirming investigation as an investigation of the actual state, important for comparing how sufficiently and credibly the investigation was in the first 3 phases and to what extent the designer used in the calculation model values very close to reality, or was too conservative, or too optimistic. this phase is typical for tunnels, also for foundation pits or for cuts of the earth structures of transport infrastructures, where the actual condition can be visually verified with the possibility of confirmation by additional tests. this possibility does not exist for the subsoil of embankments of earth structures. it is natural that for simple structures with low risk, the phasing of the gi is reduced in exceptional cases to a one-stage survey, but the "desk study" phase should always take precedence for the collection of the existing data and the considered single stage investigation should complement or specify these assumptions. 4. mutual interaction between different steps of the ground models with models of the geotechnical structure the importance of the first step, the evaluation of all available information, has already been mentioned, when for example for transport infrastructure structures, it has an influence on the choice of variants, as well as on the environmental impact assessment – eia. however, already at this stage it is possible to apply the principle of sustainable construction. two issues can be mentioned in this regard. the first one is connected with a recommendation whether the principle of reinforcement (for steeper slopes) will or will not be applied for the design of embankments and cuts. the question about required demands on land occupation and purchase become prominent as well the issue concerning volume of excavated and redeposited soil. the second is related to the possibility of using large-volume waste such as construction and demolition waste, excavated soil, and rock excavated during construction of underground structures or underground garages, etc. similarly, this applies to various wastes arising from the extraction of raw materials and their utilisation. this can significantly affect the issue of the balance of cubatures between cuts and embankments, where there is possibility of reducing the consumption of natural materials. the second step – preliminary gi focuses primarily on areas about which there is not so much prior information taken from the first step. the survey points should, in the first instance, provide sufficient information for the assessment of the individual lithological layers, when for these layers the term "geotechnical type" can be used as well. it is not only about layer interface but also about the basic properties of each layer, at least of a descriptive and index nature, when sometimes a set of results of these index properties can be used for a closer separation of individual lithological layers. this naturally also applies to information on filtration properties and groundwater levels, with respect to pore water pressures. however, with regard to the location of survey points and their depth, it is possible to use the recommendations in čsn en 1997-2 annex b.3 [5]. at this stage, however, it can only be on the basis of experience when, for example for the foundation of bridge piers, it has not yet been decided whether they can apply spread or deep foundations (piles). however, the preliminary gi should provide sufficient information for the designer’s decision about foundation engineering. as for this decision the designer can realise the preliminary design according to the limit states based on so-called table values of the mechanical-physical geotechnical parameters utilising ground index properties or different classification systems, e.g. [6]. as field investigation methods have certain advantages in this respect, there is a tendency to prioritise them when their selection for different phases of the survey is specified in annex b.2 in čsn en 1997-2 [5]. similar tables of the mechanical-physical properties of soils can be used for the preliminary design of embankment and cut slopes. for embankments, information is obtained from the places from which this material will be used, as in borrow pits of cuts of transport infrastructures. when non-standard materials are used (such as large volume waste) a previously obtained experience can be applied. preliminary design is used for the planning of the next step of geotechnical/ground investigation, as now the area affected by a proposed structure is more delineated. on the third step, the step of detailed investigation, investigation for the geotechnical structure design is crucial. additional survey points already use the preliminary design of the geotechnical structure, so they can fully follow the recommendations of čsn en 1997 with regard to location and depth. at the same time, however, when planning this phase, it is necessary to divide the individual sections of the earth structure of transport infrastructure according to the risk and the class of consequences, and into different geotechnical categories. this division into geotechnical categories thus specifies the required scope and focus of field and laboratory tests. for sections falling into the 1st geotechnical category, the design can be implemented using previous experience and so descriptive and index properties in sufficient numbers for the individual lithological layer will suffice. simply put, samples taken within the sampling category of class a and 40 vol. 32/2022 the importance of a combined 3d model the quality class of samples 2, in accordance with čsn en 1977-2, will suffice for laboratory tests. naturally, it is also possible to use mechanical-physical properties obtained for neighbouring sections of the earth structure of transport infrastructure, and coming into a higher geotechnical category for the same lithological layer. for sections falling into the 3rd geotechnical category, the gi must be planned so that for each lithological layer it is able to provide a sufficient number of samples for statistical evaluation of the measured mechanical-physical properties for these lithological layers. the number of 5 samples is usually given as a minimum for statistical evaluation. the designer then decides (and is also responsible for) the representative/ characteristic value that he/she assigns to the individual lithological layer and is thus part of the "geotechnical design model". the fourth step has the function of further specification, often coinciding with the last step, and can significantly affect the final price of the structure brought on by a change in design or in the technology of structure execution. the fifth step records the last findings on the properties of the ground/rock and soil environment and its 3d model together with a 3d model of the actual geotechnical structure or the whole structure and is the final output of the bim model and serves for subsequent decisions over time throughout the life of this structure. for transport constructions, this is a decision regarding future interaction with another construction or in necessary decision-making in the case of accidents or disasters, where there is required quick decisions on the choice of remediation method in the event of an incident involving vehicles transporting hazardous substances. 5. conclusion the interconnection of 3d models of ground and geotechnical structures, whether foundation, ground or underground structures, is today referred to as the bim model and is one of the outputs of digitisation in the construction sector. it is used for communication among partners in the gradual specification of individual phases associated with the geotechnical/ground investigation and geotechnical structure design. this eliminates possible errors, inaccuracies, and helps speed up the construction process and also defines the responsibilities of individual partners. attention is paid to the issue also because government resolution no. 82 "concept of the introduction of the bim method in the czech republic" specifies that all above-limit contracts financed from public budgets must be awarded in the bim regime from 2022. this serves to point out the importance of a 3d model of the ground connection with a 3d model of the geotechnical structure, with an emphasis on transport infrastructure. more precisely, the phasing and gradual refinement of the bim model according to the individual phases of geotechnical/ground investigation in connection with the individual phases of the geotechnical structures design has been shown. this is a preference for earth structures such as embankments and cuts and for foundation structures in the designing of transport structures. this is accomplished while respecting the principles of geotechnical structures design according to čsn en 1997 [5], in terms of risk, geotechnical categories, and in terms of requirements for the sampling class or sample quality class. references [1] i. vaníček, j. pruška, d. jirásko. bim model – application in geotechnical engineering. (in czech). in: proc. 47th conference on foundation engineering brno 2019, s. 91-98. [2] i. vaníček. sustainable construction. czech technical university press, praha, 2011. [3] i. vaníček, d. jirásko, m. vaníček. modern earth structures for transpor engineering. engineering and sustainability aspects. crc press, taylor and francis group, london, 2020, 171 p. [4] eurocode. basis of structural and geotechnical design. draft pren1990, september 2020. [5] čsn en 1997. eurocode 7 geotechnical design (in czech navrhování geotechnických konstrukcí. 1997-1, 1997-2, 2006, 2008. [6] a. bond. eurocode 7: half-term report. (invited lecture), 41st conference on foundation engineering. brno, 11th november. 41 acta polytechnica ctu proceedings 32:37–41, 2022 1 introduction 2 bim 3 models time sequence 4 mutual interaction between different steps of the ground models with models of the geotechnical structure 5 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2019.25.0012 acta polytechnica ctu proceedings 25:12–16, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app mechanical and optical investigation of laser welded structural steel poly(methyl-methacrylate) hybrid joint structures tamás csiszéra, b, ∗, tamás temesic, lászló molnára a edutus university, engineering institute, stúdium tér 1., 2800 tatabánya, hungary b óbuda university, sándor rejtő faculty of light industry and environmental protection engineering, doberdó út 6, 1034 budapest, hungary c budapest university of technology and economics, faculty of mechanical engineering, department of polymer engineering, muegyetem rakpart 3, 1111 budapest, hungary ∗ corresponding author: csiszer.tamas@edutus.hu abstract. modern welding processes that can easily be automated (such as friction stir welding, laser welding and ultrasonic welding) are gaining popularity in joining metal-polymer hybrid structures. this field of science is intensively studied around the globe, as a dependable, productive joining method that directly produces structurally sound joints between a metal and a polymer structure could unleash unforeseen possibilities in the vehicle industry. in our experiments, we manufactured hybrid steel-poly(methyl-methacrylate) (pmma) joints with laser welding, using the 2p design of experiment method. we measured the effect of cellulose reinforcing fibres (in varying weight percentages) on the transparency and weldability of the pmma material and the effect of welding parameters on the mechanical properties of the joints. we also examined the vicinity of the welded seam with scanning electron microscopy. keywords: laser welding, metal-polymer hybrid structures, poly(methyl-methacrylate). 1. introduction the importance of integrated structures made from dissimilar materials, for example by joining steel and polymer structures together is increasing in the vehicle industry. an advantage of these integrated structures is that less dense materials can be used, and thus vehicle manufacturing and operating costs can be reduced. however, the mechanical properties of the integrated structure must be comparable to the mechanical properties of the structure it replaced. besides searching for compatible metal and polymer material pairs, a considerable amount of effort is made to convert and optimise existing joining technologies that can be automatized (for example friction stir welding [1–3], ultrasonic welding [4–6], laser welding [7–10], and also a combination of laser and ultrasonic welding [11]) to make them suitable for the mass production of integrated structures made from dissimilar materials. one factor that hinders this process is the fact that metals and polymers differ from each other both in physical structure and chemical composition and thus polymers and metals behave differently when joined together [12]. if the polymer material contains additives, reinforcing fibres or nanofibers, the polymer’s behaviour during joining (for example during laser irradiation) is even more complex [13]. laser beams can be used in multiple ways during the processing and manufacturing of hybrid or integrated metal-polymer structures: depending on the laser welding machine and other parameters like laser power and beam velocity, laser beams can be used to cut materials, to process and to prepare their surfaces and also to join them together. the latter is usually called laser assisted metal polymer, or lamp joining for short in publications [14]. the possibility of using laser welding to join metal and polymer structures together is already proven in publications [15]. rauschenberger et al. [8] proved that using the ideal process parameters during the joining of steel and polymer specimens (both reinforced and unreinforced) can result in even better joint strength than the strength of the polymer materials themselves. jung et al. [16] joined zinc-coated steel and carbon fibre reinforced polyamide 6: the ultimate shear force of the overlapped joint reached 3300 n. based on transmission electron microscopy (tem) and energy dispersive x-ray spectroscopy (eds), the authors claimed that the strength of the macroscopic joint was influenced by shape-connected mechanical joints: the molten polymer filled in the surface roughness grooves of the zinc coating. they also found that chemical bonding between zinc-oxide molecules and carbon atoms of the polymer chains further increased the joint strength. it is also stated by cheon et al. [17] that the crystalline structure, the primary chemical bonding structure and the conformation of the polymer chains have a significant effect on the strength of metal-polymer joints. bubbles can form in the polymer material during the joining process caused by the evaporation of remnant monomer molecules or 12 http://dx.doi.org/10.14311/app.2019.25.0012 http://ojs.cvut.cz/ojs/index.php/app vol. 25/2019 mech. and opt. investigation of laser-welded steel-pmma hybrid joints additives (this phenomenon is called humping), which can strengthen or weaken joints depending on their quantity and size [18]. adhesion between the metal and polymer joining partners can be increased by structuring the joining surface of the metal specimen using a laser beam either to remove material from the surface [19], to transform material on the surface [20, 21], or to add material to the surface [5, 22]. it is also possible to use technologies like corona-discharge plasma treatment to chemically modify the surface of either the metal and/or the polymer specimens. in our ongoing research, we joined structural steel to unreinforced and reinforced (with bio-based cellulose fibres in varying weight percentages) poly(methylmethacrylate) (pmma) specimens using two different laser welding machines. our aim was to identify and optimise the technological parameters that had a significant effect on the strength and appearance on the joint and to find and adjust measurement methods that are suitable for the measurement of the joints we made. 2. materials and methods 2.1. materials and specimens materials used in this work were structural steel (s235jr), 50×10×0.8 mm in dimension as metal specimen, and cellulose fibre-reinforced (in 0 . . . 10 wt% ratio, whole numbers only) pmma, 80 × 10 × 2 mm in dimension made by two manufacturers: • the first pmma type we used is sold under the trade name sitramac hw55, a pmma comonomerised with polystyrene (ps) made by sitraplas gmbh, • the second pmma type we used is sold under trade name altuglas vs-uvt, a pure and crystal-clear pmma made by the arkema group. firstly, we measured the effect of cellulose fibres on the transparency and weldability of the pmma material. for this, we manufactured cellulose fibre-reinforced specimens using the sitramac base material and cellulose fibres manufactured by kronospan-mofa hungary ltd., using cellulose from multiple tree species. both the pmma pellets and the cellulose fibres were dried in a drying oven for 8 hours on 80 °c (as specified by the manufacturers) before compounding on a twin-screw extruder (labtech lte 26-44, ascending temperature profile with 5 °c increments, temperature at the rod-type die: 200 °c). an endless filament was formed from the molten compounded material, which was pelletized for injection moulding on an arburg allrounder 270s 400-170 injection moulding machine. standard (dogbone-shaped) tensile test specimens were manufactured. later on, we manufactured further reinforced and unreinforced specimens for testing and joining using both the sitramac and the altuglas vs-uvt base materials. the reinforcement we used was 1 wt% cellulose fibre (manufactured by arbocel under the trade name b600) with the same process parameters on the same twin-screw extruder and injection moulding machine. flat, rectangular specimens 80 × 80 × 2 mm in dimension were injection moulded, from which, 80 × 10 × 2 mm specimens were cut using a disk cutting machine for joining. 2.2. experimental methods as stated in the introduction section, the chemical composition, material structure of the polymer and the reinforcing material can influence the behaviour of the specimen during laser welding, so we measured the transparency of the specimens to visible and nearinfrared (nir) light on a perkin-elmer lambda 1050 spectrophotometer. we found, that in the 950-1050 nanometre range (in which our laser welding machines operate), the direct transparency (the amount of light passing through the material without scattering) of the unreinforced specimens reached an average of 75 % and 82 % (for sitramac and altuglas base materials, respectively). when cellulose fibres were added to the base material in 1 wt% ratio, the direct transparency of the specimens fell to an average of 35 % and 22 %, respectively. adding more than 1 wt% cellulose fibres to the base material decreased the transparency so much that basically all the incident light scattered or were absorbed inside the material (see figure 1). we used two different laser welding machines to create two different kinds of overlapped joints between the steel and polymer specimens. all metal and polymer specimens were degreased and cleaned using methanol before joining. the first type of the overlapped joints was manufactured with the so-called transmission welding process, in which the polymer specimen is placed on top of the steel specimen. in this case, the laser beam passes through the polymer and then it is absorbed in the steel specimen, creating heat. the heat is then conducted to the polymer specimen, which in turn melts and fills in the surface roughness grooves of the steel, forming shape-connected (mechanical) joints. the second type of process we used was the so-called direct laser welding. in this process, the polymer specimen is placed below the steel specimen, while the latter is irradiated and heated up by the laser beam. with this method, higher laser power outputs may be used and thus, faster cycle times can be achieved without the degradation of the polymer material. we used a trumpf trudiode 151 diode laser with a maximum power output of 150 w for the transmission welding process, and a trumpf trudisk 4001 (yb:yag) solid state laser with a maximum power output of 4000 w for the direct laser welding process. the main process control parameter was the laser power. during the transmission welding with the diode laser, laser power was set to 30, 40 and 50 w , while during the direct laser welding with the solid 13 t. csiszér, t. temesi, l. molnár acta polytechnica ctu proceedings figure 1. direct transparency of unreinforced and reinforced pmma specimens rate meaning 0 no data 1 no flame 2 small flame 3 moderate flame 4 heavy flame table 1. the qualification of flaming rates state laser, laser power was set to 210, 230 and 250 w . these values were chosen based on preliminary measurements we conducted. we used the 2p design of experiment method to manufacture a total of 200 steelpmma specimens. all other parameters remained constant during the joining process: the specimens were joined in 3 cycles in a straight, 10 mm line, while the welding speed was set to 1 m/min. investigating the mechanical (shear) strength of the joined specimens is ongoing. in the next section, we will present a qualitative analysis of the joined specimens and some preliminary results. 3. discussion in most cases during the direct welding process, the polymer specimens burst into flames. the degree of flaming was qualified according to its intensity (see table 1) with visual inspection during the welding process. applying 250 w of laser power, the average rates of intensity of flaming were between 3 and 3.5 for each matrix-fibre combination, which proves the presumption that higher energy input can cause the intensive degradation of the polymer, which can lead to flaming. on lower laser power levels (at 210 and 230 w ), the flame intensity of the materials were similar, except for the sitramac pmma reinforced with 1 wt% cellufigure 2. average flaming rates during direct laser welding of steel-pmma specimens lose fibres at 230 w laser power (figure 2). a possible reason for this phenomenon can be that the chemical composition of the sitramac pmma differs from that of the altuglas pmma. based on the tensile shear testing of overlapped joints made by the solid-state laser with 250 w laser power between unreinforced (sitramac) pmma and steel, we can say that the joints are quite rigid and standard deviations of both ultimate shear force and displacement values are high (figure 3). this may be caused by inadequate contact (air gap) between specimens during joining, or the fact that the material surfaces were only degreased before the joining process. in later experiments, we made sure that there wasn’t an air gap between specimens during joining, and we are planning on trying different techniques to prepare the surfaces of the specimens in the hopes of reaching better shear strength values. we also noticed that after the tensile shear tests, some pmma remained on the steel specimen’s surface. using a scanning electron microscope (sem), we discovered that the connection between the steel and pmma was not evenly distributed on the join14 vol. 25/2019 mech. and opt. investigation of laser-welded steel-pmma hybrid joints figure 3. ultimate tensile-shear force values of unreinforced pmma-steel specimens figure 4. cohesive failure of the pmma near the steel-polymer interface (sem image of the steel surface) ing surface. in some areas, the adhesion between the materials was so good, that the failure was inside the pmma material (as shown in figure 4, where the plastic deformation of the pmma material can be seen), while in other areas, the joint failed with simple adhesive failure (debonding) on the interface. the pmma material also possibly degraded: the formation of gas bubbles on and near the joint interface can clearly be seen in figure 5, however, at lower laser power values, no joints could be manufactured. the degree of degradation of the pmma and the possibility of forming chemical bonds between the materials needs to be investigated further. 4. conclusions with our experiments, we proved that it is possible to manufacture joints between pmma and steel specimens with laser welding equipment. both the transmission welding (when the laser beam passes through the polymer material) and the direct laser welding (when the laser beam does not pass through figure 5. gas bubbles on and near the steel-polymer interface and adhesive failure of the pmma (sem image of the pmma surface) the polymer) techniques could be used to join the specimens. we also proved that cellulose fibres greatly decreased the direct transparency of the otherwise crystal clear pmma material even at low weight percent ratios, thus making the transmission welding joining of reinforced pmma and steel specimens difficult and time-consuming. the steel-polymer joints we manufactured were rigid and the standard deviation of measured tensile force values were high, which was possibly caused by the degradation of the pmma material during welding and the uneven connection between pmma and steel in the joint area. in most cases during the direct welding process, the polymer specimens burst into flames, which also decreased joint strength.the two prominent failure modes were adhesive failure (debonding) on the joint interface and cohesive failure of 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[22] d. p. graham, a. rezai, d. baker, et al. the development and scalability of a high strength, damage tolerant, hybrid joining scheme for composite-metal structures. composites part a: applied science and manufacturing 64:11–24, 2014. doi:10.1016/j.compositesa.2014.04.018. 16 https://doi.org/10.1016/j.matchar.2014.10.013 https://doi.org/10.1080/13621718.2016.1204171 https://doi.org/10.1007/s40194-017-0517-y https://doi.org/10.1002/adem.200800271 https://doi.org/10.1016/j.compstruct.2014.07.019 https://doi.org/10.1016/j.ijadhadh.2015.11.001 https://doi.org/10.1016/j.scriptamat.2008.08.026 https://doi.org/10.1117/12.2080226 https://doi.org/10.1016/j.optlastec.2016.09.028 https://doi.org/10.1016/j.proeng.2017.04.092 https://doi.org/10.1016/j.jmatprotec.2017.06.036 https://doi.org/10.1002/9781119429807 https://doi.org/10.1016/j.matdes.2016.03.044 https://doi.org/10.1016/j.matdes.2012.12.015 https://doi.org/10.11779/1362171814y.0000000236 https://doi.org/10.1016/j.phpro.2011.03.030 https://doi.org/10.1016/j.jmatprotec.2015.10.026 https://doi.org/10.1016/j.phpro.2011.03.065 https://doi.org/10.1007/s00170-016-9646-5 https://doi.org/10.1016/j.compositesa.2014.04.018 acta polytechnica ctu proceedings 25:12–16, 2019 1 introduction 2 materials and methods 2.1 materials and specimens 2.2 experimental methods 3 discussion 4 conclusions 5 acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.29.0001 acta polytechnica ctu proceedings 29:1–5, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app the oldest underground railways in the world and their anniversaries jiří barták czech technical university in prague, faculty of civil engineering, department of geotechnics, thákurova 7, 166 29 praha 6, czech republic correspondence: bartakj@fsv.cvut.cz abstract. the five oldest underground railways started to be operational in the 19th century. these were the metros in london, athens, istanbul, budapest and glasgow. undoubtedly, world primacy of status in terms of age belongs to the london underground. however, the age ranking of the others in continental europe is more difficult to determine. regardless of this criteria, the undergrounds of the 19th century became the basis of the world’s most widespread and rational means of public transport. the further development of these named oldest undergrounds is here documented. keywords: metro anniversary, metro establishment, underground railway. 1. introduction the nineteenth century is often characterised, somewhat figuratively but justifiably, as a century of steam. in 1825, george stephenson, an english mechanical engineer, started the public operation of a train pulled by a steam locomotive along a track between stockton and darlington. his locomotive, rocket, and newly invented hot rolled rails passed successfully into technical operation. in 1830, an indication of the future use of the underground for transportation purposes appeared on the manchester – liverpool rail track which was the first railway tunnel on a track using steam trains. this was the 1,066m long edge hill tunnel. numerous fundamental things required for the commencement of operations in the underground (steam locomotives with the necessary power, railbound passenger cars, rolled rails and their fixing, design of trackwork and other technical equipment) were tried on railway tracks. thus, the conditions for the creation of an efficient urban mass transit system – the underground railway – had been prepared. world leadership in the development of the highestcapacity and most widely spread urban mass transport is undisputable – the first six kilometres long line of the london underground went into service on 10th january 1863. on continental europe however there is a certain problem with the determination of the first metro construction project – there are three candidates. arranged chronologically, they are the metros in athens (1869), istanbul (1874) and budapest (1896). the instigation and subsequent development of metro networks in those cities exhibits noteworthy particularities and differences, making it relatively difficult to award the second place to one of them unambiguously. nevertheless, it is a straightforward fact that celebrations of the original creation of the first metro in continental europe are held in all of the three countries. 2. london underground the first great exhibition was held at crystal palace in london in 1851. the possibility of developing an underground railway was first mentioned at the exhibition. the preparatory work continued, and a company commercially registered under the name of metropolitan commenced constructing the london underground by connecting the end stations of two railway lines – the great western railway and the great northern railway. the basis of the london underground was formed by developing a 6km long section between paddington station and the farringdon street railway station ([1], [2], [3]). with respect to an operation using steam engines and the limited capacity of ventilation through chimneys, the grade lines of tunnels were designed to be at small depths (4 to 8 metres). horseshoe-shaped profile double-track tunnel linings were constructed on wood centrings in pits dug in the space of streets and were backfilled after completion (see fig. 1). figure 1. historic subsurface tunnels of the london underground [3] the first line in the world using wooden cars pulled 1 http://dx.doi.org/10.14311/app.2020.29.0001 http://ojs.cvut.cz/ojs/index.php/app jiří barták acta polytechnica ctu proceedings by a locomotive was brought into service with great success in london at the beginning of 1863 (10th january 2018). the first day of operation was attended by 38,000 visitors, and later it was used on average by 25,000 passengers every day. given the extraordinary attendance rates and the interest of citizens in this new mode of transport, this subsurface variant of the london underground using steam engines was considered for further extension. however, it was only after the introduction of electrification into the underground in 1890 and the use of an innovated tunnelling shield with a circular profile that it became possible to successfully drive couples of single-track tunnel tubes with a diameter 3,6m through deeper parts of london. the connection of the 1st line of the london underground with the legendary thames tunnel is very interesting. the thames tunnel was constructed under the river in muddy and sandy alluvium by civil engineer marc isambard brunel, who, in 1918, took out a patent for a completely new tunnelling method of excavation and supporting the excavated space by means of the world’s first tunnelling shield, which became much later on the basis of tunnel excavation technologies using tunnel boring machines for all types of rock and soils. the thames tunnel, which was awarded the status of a technical monument, was meticulously reconstructed and, since 2010, has become part of the london overground railway network, which is an extensive system of commuter railways serving a significant part of london and its regional environs. the current london underground has 11 lines with 292 stations. with its length of circa 408km it was the world’s longest underground railway system until 2012. trains corresponding to common railway transport services are operated on the subsurface lines; specially adapted trains are used on deep subsurface lines, in single-tube tunnels with the diameter of only 3.6m (see fig. 2). outside of central london, the underground trains run along deep-seated lines at speeds up to 96km (60miles) per hour. figure 2. cars for subsurface and deep-seated london underground lines [3] 3. the metro in athens in 1969, the 100th anniversary of opening the second underground railway in europe and also the first one in the continental part of europe, (connecting the port of piraeus with the greek capital), was celebrated in athens. the circumstances of the instigation and development of this line have often raised doubts over whether it is really a metro; nevertheless, greeks (and many other people) consider it to be the fully-fledged germinal source of the later athens metro system. at about 10km long, at-grade single-track railway line was constructed during the course of two years on the basis of a contract with edward pickering, an english businessman ([2], [4]). it was inaugurated on 27th february 1869. it connected the railway station in piraeus with the thiseio railway station in athens. during the ceremonial beginning of this service supported by the state, and attended by queen olga and the prime minister, a steam locomotive with a train set consisting of ten cars covered the distance between the two cities in 15 minutes. the line was extended to the inner city, omonia square, by a subsurface section in 1893, and was electrified in 1904. the at-grade portion stretch of the line still predominated. the line waited for subsequent changes until 1930 (doubling of the track – see fig. 3), later for the extension of the track to attiki station in 1948; and the whole route up to the kifisia end station was brought into service in 1957 as the first metro line. the length of this single line 1 route amounts to 25.6km, with a 3km long section running underground. only after a subsequent 54 years (in 2011), was line 1 incorporated into the athens underground railway system (the 150th anniversary of the commencement of operation will be celebrated in 2019). figure 3. original athens metro line after doubling of the track [4] currently athens metro (aττικó mετρó ) has 3 lines (see fig. 4) with an aggregate length of 84.5km. the work on the underground lines 2 and 3 commenced as late as 1991. the initial sections were brought into service in 2000 and were further extended; the current length of the lines amounts to 58.9km. the construction of a line 4 is being planned. it should have a circular character. line 1 with its historic piraeus – 2 vol. 29/2020 the oldest underground railways in the world thiseio – omonia square section was operated until 2011 separately from the newly developed metro system; now it is a part of it. 4. the metro in istanbul the only underground line was inaugurated on 23rd september 1875. it is 573m long. the 140th anniversary of bringing the second metro in continental europe was held in turkey in 2015 ([5], [4]). however, it raises rather legitimate doubts whether it can be considered strictly to be a metro. in this case it is obvious that it is not really the original start point of the future metro because of the fact that it has not been incorporated into the modern istanbul metro system. nevertheless, one primacy of status cannot be denied to this “metro” – it is the shortest one in the world. this solitary underground line, which is currently known by the name of tünel, was designed by henri gavand to be provided with a rope drive. it allowed for a more more comfortable overcoming of the sixtymetre difference in elevation with a steep rise of the surface between karaköy and beyoğlu, two city districts lying at the european corner of a bay on the aegean sea. the line has only two stations, boarding (see fig. 4) and disembarking, both with side platforms. a double-car train set covers the distance between the stations in 1.5 minutes, running at a speed of circa 25km per hour. the single-rail track has a passing bay in the middle of the route. the car undercarriages are on rubber tyres and move along a concrete track. figure 4. karaköy station of istambul metro [3] in 1910, the line was electrified and between 1968 and 1971 it was significantly modernised. the line has been incorporated into the istanbul at-grade transportation system. it links to a common tramway track at the boarding station (itself linking to the karaköy port); the disembarking station links to the historic tramway track (transfer to line m2 of the classical metro being also possible). the line carries a daily average of 15,000 passengers. the current istanbul metro (istanbul metrosu), which started operating on 3rd september 1989, had eight lines with an aggregate length of 105.8km and 82 stations in 2018. there are 6 classical metro lines marked as m1 through to m6 and two lines of the light rail transit system marked as t1 and t2 also in operation. the system is undergoing substantial further extensive development – 5 additional lines with 62 stations are under construction. the metro is connected with the taksim-kabataş underground funicular (2006) and the trans-bosporus marmaray undersea railway tunnel (2013). 5. the metro in budapest while the budapest metro ([2], [6], [4]) was brought into service as late as 2nd may 1896, it must be considered that, in terms of its actual attributes, it unambiguously meets the qualification requirements for second place among the european metros, compared to the earlier operating structures in athens and istanbul. it was inaugurated 33 years after the london underground. in 2016, the 120th anniversary of the metro was celebrated in hungary. in 1896, the volume of traffic under andrassy avenue, the main link of central pest to the rapidly growing new residential development more remote from the left bank of the danube, reached an unacceptable level. the solution was found in constructing an underground rail line which was, with respect to its character, rightly named as kisföldalatti (ie small underground space), and with time this became abbreviated simply to földalatti). the commencement of construction work and its rapid continuation was probably significantly affected by a more obviously political and historic aspect – the construction of the metro and the bringing of it into service was simultaneously conceived as a celebration of the thousand-year anniversary of the arrival of hungarians to pannonia. the contract for the construction was awarded to the budapest railway company under the condition that it would be completed before the millenium celebrations. the railway company fulfilled the condition and the so-called jubilee line, 3.7km long with 11 stations, was brought into service in may 1896, only 20 months after the commencement of the construction work. virtually the whole line, with the exception of a short at-grade section at its end, was constructed in sloped or braced pits (see fig. 5), fast completion being therefore necessary even from the viewpoint of surface transportation. the framework structure of the reinforced concrete tunnel lining was minimised to the dimension of circa 3 × 4m, and the height of the overhead catenary was only 2.3m above the top of the rail. the interior design of the stations was subject to quite demanding conditions set by the budapest city hall. the art nouveau style was chosen. the walls of the stations were clad with white and black tiles, while cast-iron columns along the edge of the short platform (circa 40m long) were provided with decorative heads, 3 jiří barták acta polytechnica ctu proceedings figure 5. földalatti under andrassy avenue [6] figure 6. kossuth lajos square five-vault station on line m2 [2] rest benches and cashier kiosks designed in the same style (see fig. 6). the structure of entrance staircases leading to stations was noteworthy – the treads were unusually wide, and risers were low (even entering on horseback was reportedly possible for officers). metro cars were wooden; before their modification for the underground they were originally intended for tramway services. the first very necessary modification took place as late as the 1930s, when, in particular, the power of engines was increased, doors of the cars were reconstructed and the traction voltage was changed. however, the belgian-produced wagons were operated until the 1960s. the földalatti line is newly identified as m1. it remained a sole provision in budapest until 1970, when line m2 of the modern metro was brought into service. the current budapest metro (budapest metró) has experienced an interesting evolution. the construction of line m2 commenced in 1950; however, the work was suspended after three years and the building capacities were shifted in the direction of new residential development. until 1963, the development of the metro was discontinued, and the finished tunnels and stations were used as stores. after the resumption of the works, line m2 was inaugurated on 2nd figure 7. kossuth lajos square five-vault station on line m2 [2] april 1970. the stations on line m2 partially copied soviet-type 3-vault stations. but a new concept of budapest-type stations was also instigated (fiveto six-vault stations – see fig. 7). the budapest metro network is currently formed by 4 lines (m1 through to m4) with an aggregate length of 40km and 52 stations; the development of a line m5 is now under preparation. 6. the metro in glasgow the fourth metro, the last one brought into service in the 19th century, is the metro in glasgow, and the second on the british isles ([1], [4]). it went into service on 14th december 1896, in the same year as the metro in budapest, but seven months later. in terms of the actual system, the glasgow metro is a certain curiosity, but there are no doubts that it belongs to the category of underground railways. it is formed by only one 10.5km long orbital line, which was never expanded during following years. projects to expand it existed in several versions but none of them was realised, probably with respect to extensive mine galleries (in large part inundated) under the city surface. the glasgow orbital metro (officially named the glasgow subway, in guidebooks even referred to by a nickname „clockwork orange“, see fig. 9) is a double-track system, with each track running through a separate tunnel tube. the diameter of the tubes is probably unique in the world – only 3.35m (11 feet); the track gauge is also atypical – 1,219 mm (4 feet). figure 8. glasgow orbital metro system [4] in terms of technology, the metro was designed in a very original way. trains were hauled by endless cables powered from a steam power plant. on the inner circle, 4 vol. 29/2020 the oldest underground railways in the world the car sets moved in an anti-clockwise direction; on the outer circle, they moved in the opposite direction. all 15 stations were solely equipped with intermediate platforms. for that reason the metro cars had doors only on one side. there were no cross-over tunnels with switches between the circles; no pull cable for traversing the cars to a depot was installed. the cars had to be lifted from the metro route with a crane to be carried to the depot. the cable drive was removed as late as 1935, when the metro was electrified. current glasgow metro (glasgow subway) is, in terms of the extent, the same as it was 123 years ago – a length of 10.5km and 15 stations. in 1977, it was significantly modernised, equipped with new car sets, a connection with the new depot was installed and electrification of the system was modernised. the minimum diameter of the tunnel tubes (see fig. 10) and the reduced track gauge was preserved. a welldeveloped proposal for metro expansion, using the already mentioned old mine galleries under the city surface and a relic of historic coal mining, has existed since 2005. figure 9. current glasgow subway – kelvinhall station [4] in 2016, the glasgow transport agency issued a report that the metro operation system would be reconstructed in the coming years to an unmanned system with the cbtc (communication based train control) system of traffic control and signalling. 7. conclusion all five structures mentioned in this paper originated in the second half of the 19th century. by their mostly successful operation, they clearly showed the path that most rational solutions to urban mass transit in large metropolises would take – underground railways. despite the fact that the development of these oldest five metros was significantly variable, the metros have usually become bases of extensive networks of underground lines (london, athens, budapest) or have initiated further development of the metro (istanbul), or exceptionally the original arrangement of the metro has remained as a final solution until the present day (glasgow). it is necessary for the sake of fair and complete information to mention the urban underground railway in wien (vienna) [4], where operation actually commenced in 1898 even though it did not have the official status of a metro. a decision was made in the 1960s that the obsolete and unsatisfactory urban railway system would be reconstructed to metro (u-bahn) and a rapid transit system (s-bahn). the reconstruction of this already inadequate urban railway system into a metro system then commenced in 1969. references [1] j. streit. tunely všech dob a světadílů. publishing house karel synek, 1. edition, praha 1947. 130 pages. [2] j. křivánek, z. šmíd, j vítek. všechna metra světa. nakladatelství dopravy a spojů, 1. edition, praha 1986. 208 pages. [3] london. https://en.m.wikipedia.org/wiki/london. [4] metro. https://cs.m.wikipedia.org/wiki/metro. [5] j. straka. podzemní stavby. nakladatelství technické literatury, 1. edition, praha 1966. 442 pages. [6] u-bahn museum budapest. universitätsdruckerei, budapest 1977. 40 pages. [7] https://wikimediafoundation.org. 5 https://en.m.wikipedia.org/wiki/london https://cs.m.wikipedia.org/wiki/metro https://wikimediafoundation.org acta polytechnica ctu proceedings 29:1–5, 2020 1 introduction 2 london underground 3 the metro in athens 4 the metro in istanbul 5 the metro in budapest 6 the metro in glasgow 7 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0030 acta polytechnica ctu proceedings 26:30–33, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app mechanical properties of cement composite containing recycled concrete and alkali activated blast furnace slag jan horych∗, pavel tesárek, zdeněk prošek czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jan.horych@fsv.cvut.cz abstract. recycling of materials is very popular and very important in these days. finding the new ways to process and use these materials is a key to get rid of a lot of construction waste. the amount of landfilling needs to be reduced. this study observes mechanical properties of the cement composites containing recycled concrete powder and alkali-activated blast furnace slag processed on a high-speed mill as a potential binder replacement up to the 60 wt. %. these materials have a positive effect on hydration process, increase flexural strength. it can reduce compressive strength loss when an amount of cement in the mixture is reduced. keywords: composites, waste, blast furnace slag, recycled concrete, cement matrix. 1. introduction according to the current recycling trend, we want to find a way to use waste material as a material to create something new. in this research we use recycled concrete and blast furnace slag to replace cement and reduce impact of its production. cement is the most important but the most expensive ingredient in concrete. the production of cement is based on natural resources and its processing needs a lot of energy. unfortunately, these used materials have their own issues and problems that have to be solved. many studies are focused on similar topic [1–6]. some of these studies are using blast furnace slag or recycled concrete to improve mechanical properties. the canadian study [7] is similar to our research. it uses the blast furnace slag as a binder and examines all physical properties of concrete including workability, air content, elastic modulus, strength, shrinkage, durability and carbonation. the results of all test confirmed the opinion, that the blast furnace slag can be used as a binder. it has a positive effect on mechanical properties and durability of concrete. another study [8] solves the issue of recycled concrete powder used as a binder. they also used the high-speed mill to activate the material. they created samples with varying cement and recycled concrete percentages and observed mechanical properties of the samples. results of this study showed the possibility to use recycled concrete in practise. concrete with cement and recycled concrete produce higher hydration heat. it is caused by unhydrated cement grains contained in recycled concrete powder. research [9] was based on fly ash, blast furnace slag and recycled concrete aggregate. the study investigated compressive and tensile strengths, elastic modulus, coefficient of permeability and durability in chloride and sulphate solutions. blast furnace slag had positive effect for a concrete with recycled aggregate and could reduce strength losses. fly ash had worse results and was recommended for further testing before being used in practice. the study [10] was focused on creation of alkali-activated foamed concrete with part of ground granulated blast furnace slag. all samples were produced using a pre-formed foam with no filler or aggregate. the results revealed that foamed concrete with blast furnace slag has a potential to be used in practise. it had even better mechanical properties than foamed concrete only with portland cement. 2. materials and samples testing samples were composed of portland cement cem i 42.5 r from radotín. another component was the recycled concrete micronized by high-speed mill in collaboration with lavaris corporation. the source of concrete is 100 years old construction concrete from the demolished building on waltrovka in prague. another component was the blast furnace slag kept from kladno. it is 100 years old air-cooled and carbonated slag stored on a heap and it’s the waste from slag aggregate production. this kind of slag is not suitable as a concrete binder. it needs to be activated. we use the same high-speed mill to activate it, because the blast furnace slag had a slow reaction for a long period of time. all of the samples were 40 × 40 × 160 mm and they were stored in water for 28 days. during this time we made tests to find dynamic modulus of elasticity using a non-destructive resonance method and at the end, we made destructive bending test and compresion test to figure out final strength. samples contain 50 or 40 wt. % of portland cement, 50 wt. % of recycled concrete and changing share of blast furnace slag and slaked lime from 0 to 10 wt. % (ta30 https://doi.org/10.14311/app.2020.26.0030 https://ojs.cvut.cz/ojs/index.php/app vol. 26/2020 mechanical properties of cement composite ... mixtures portland recycled slaked blast furnace flow cement [wt. %] concrete [wt. %] lime [wt. %] slag [wt. %] expansion test [cm] a 50.0 50.0 0.0 0.0 22.0 b 40.0 50.0 0.0 10.0 21.2 c 40.0 50.0 2.5 7.5 19.0 d 40.0 50.0 5.0 5.0 19.0 e 40.0 50.0 7.5 2.5 18.0 f 40.0 50.0 10.0 0.0 17.3 table 1. composition of individual samples. ble 1). it also shows the results of flow expansiom test. the part of slaked lime changes the result even when the water coefficient w/b is the same. the slaked lime is very finely ground and because of this fact, the mixture has a different consistency. the water coefficient w/b is set to 0.4. figure 1. final bulk density of the samples. figure 2. development of the dynamic young’s modulus for the samples. figure 3. development of the dynamic shear modulus for the samples. figure 4. flexural strength of the samples. figure 5. compressive strength of the samples. 3. measurement methods during our research we measured dynamic young’s modulus, dynamic shear modulus, flexural and compressive strength. dynamic young’s modulus and dynamic shear modulus were monitored using nondestructive resonance method during 28 days. the method is based on measuring the natural frequencies of longitudinal, transverse and torsional vibration of the samples. samples were supported with a soft elastic pads at nodal points of basic natural mode shapes. through the test we used brüel & kjær measurement station type 3560-b-x12 to record the excitation and response signals. as an acceleration transducer was used brüel & kjær type 4519-003 and another component is brüel & kjær impact hammer type 8206. everything was connected to the computer with software pulse labshop. the excitation and response signals were transformed from the time domain to the 31 j. horych, p. tesárek, z. prošek acta polytechnica ctu proceedings frequency domain by using the fast fourier transform. frequency response function was calculated as a ratio of the response and the excitation force. the basic natural frequencies were evaluated with frequency response function and then were evaluated dynamic young’s modulus and dynamic shear modulus. in this method we also need to know the dimensions and the weight of the samples. [11, 12]. flexural strength was measured on 28 days old samples using three-point bending test. the testing was displacement controlled at a constant rate of 1 mm/min and the distance between supports of the sample was 100 mm. during the test, samples were broken approximately in half. these halves (40 × 40× ∼ 80 mm) were used on the uniaxial compression test. the testing was displacement controlled at a rate of 5 mm/min [13]. 4. results and discussion figure 1 shows the final bulk density of all samples. the highest bulk density had the sample a as a reference. it had 50 wt. % of portland cement and 50 wt. % of recycled concrete powder. the second highest bulk density had the sample c with 40 wt. % of portland cement, 50 wt. % of recycled concrete powder, 7.5 wt. % of blast furnace slag and 2.5 wt. % of slaked lime as the activator. as we can see in figure 2 and figure 3, sample c had also the highest dynamic young’s modulus and the dynamic shear modulus from the samples with 40 wt. % of portland cement. however the reference sample had the higher dynamic young’s modulus approximately by 3 gpa and dynamic shear modulus by 1 gpa. each sample was measured during 28 days every 7 days including its weight and dimensions. in both cases, dynamic young’s modulus and dynamic shear modulus had the highest progress in the first 7 days. then the progress was much slower from day 7 to day 21 and in the last 7 days, the values were almost unchanged. figure 4 presents a flexural strength results of all samples. it is obvious that the blast furnace slag had a positive effect on flexural strength. the worst strength had a reference sample a without the slag. samples b, c and d had a higher strength almost by 100 %. however, the highest strength had the sample c and confirmed the previous measurement. figure 5 shows the results of the compression test. it proved the effect of recycled concrete and slag. the highest compressive strength had a sample a because of the highest amount of cement in the mixture. as we can see, the samples c, d and e had approximately the same strength. the combination of blast furnace slag and slaked lime seemed to be effective. through all stages of testing, the results seemed very promisinig. it seemed that the most promising sample was sample c because of the highest flexural strength. compressive strength was also one of the highest, so the ratio 3:1 of the slag and slaked lime seemed to be ideal. recycled concrete could be used as a microfiller but, the study [6] proved that very fine ground concrete has approximately 10 % of unhydrated cement grains and these grains helps with hydration process. there is a need for more research, but used materials definitely have the potential to be used in practical applications. 5. conclusion this paper examined the recycled concrete powder and blast furnace slag as a cement replacement at the level of 60 wt. % with various ratios of blast furnace slag and slaked lime and its effect on mechanical properties of samples. based on the results, it can be concluded that: • amount of cement in concrete can be reduced without a massive loss of mechanical properties. • highly carbonated blast furnace slag can be used as a filler and also a binder with proper processing. • blast furnace slag had a very positive effect on flexural strength. • the ratio 3:1 of blast furnace slag and slaked lime seems to be ideal. acknowledgements this outcome was supported by the czech technical university in prague under no. sgs19/148/ohk1/3t/11 and by technology agency of the czech republic (ta cr) research project tj01000435, which is solved in cooperation with the company pozemní komunikace bohemia, a.s.. the authors also thank the center for nanotechnology in civil engineering at the faculty of civil engineering, czech technical university in prague. references [1] s. demirdag, l. gunduz. strength properties of volcanic slag aggregate lightweight concrete for high performance masonry units. construction and building materials 22(3):135–142, 2008. doi:10.1016/j.conbuildmat.2006.10.002. [2] j. sun, j. feng, z. chen. effect of ferronickel slag as fine aggregate on properties of concrete. construction and building materials 206:201–209, 2019. doi:10.1016/j.conbuildmat.2019.01.187. [3] r. arellano aguilar, o. burciaga díaz, j. escalante garcía. lightweight concretes of activated metakaolinfly ash binders, with blast furnace slag aggregates. construction and building materials 24(7):1166–1175, 2010. doi:10.1016/j.conbuildmat.2009.12.024. [4] j. schoon, k. de buysser, i. van driessche, n. de belie. fines extracted from recycled concrete as alternative raw material for portland cement clinker production. cement and concrete composites 58:70–80, 2015. doi:10.1016/j.cemconcomp.2015.01.003. [5] j. schoon, l. van der heyden, p. eloy, et al. waste fibrecement: an interesting alternative raw material for a sustainable portland clinker production. construction and building materials 36:391–403, 2012. doi:10.1016/j.conbuildmat.2012.04.095. 32 https://doi.org/10.1016/j.conbuildmat.2006.10.002 https://doi.org/10.1016/j.conbuildmat.2019.01.187 https://doi.org/10.1016/j.conbuildmat.2009.12.024 https://doi.org/10.1016/j.cemconcomp.2015.01.003 https://doi.org/10.1016/j.conbuildmat.2012.04.095 vol. 26/2020 mechanical properties of cement composite ... [6] z. prošek, j. trejbal, j. topič, et al. utilization of the waste from the marble industry for application in transport infrastructure: mechanical properties of cement pastes. iop conference series: materials science and engineering 236(1), 2017. doi:10.1088/1757-899x/236/1/012092. [7] r. d. hooton. canadian use of ground granulated blast-furnace slag as a supplementary cementing material for enhanced performance of concrete. canadian journal of civil engineering 27(4):754–760, 2011. doi:10.1139/l00-014. [8] j. topič, z. prošek. properties and microstructure of cement paste including recycled concrete powder. acta polytechnica 57(1):49–57, 2017. doi:10.14311/ap.2017.57.0049. [9] m. l. berndt. properties of sustainable concrete containing fly ash, slag and recycled concrete aggregate. construction and building materials 23(7):2606–2613, 2009. doi:10.1016/j.conbuildmat.2009.02.011. [10] k.-h. yang, k.-h. lee, j.-k. song, m.-h. gong. properties and sustainability of alkali-activated slag foamed concrete. journal of cleaner production 68:226– 233, 2014. doi:10.1016/j.jclepro.2013.12.068. [11] astm 1876-01. standard test method for dynamic young’s modulus, shear modulus, and poisson’s ratio by impulse excitation of vibration, annual book of astm standards, american society for testing and materials 2006. [12] čsn 73 1372. in nedestruktivní zkoušení betonu rezonanční metoda zkoušení betonu, úřad pro technickou normalizaci, metrologii a státní zkuševnictví,. 2012. [13] čsn en 12390 (731302). in zkoužení ztvrdlého betonu. 2013. 33 https://doi.org/10.1088/1757-899x/236/1/012092 https://doi.org/10.1139/l00-014 https://doi.org/10.14311/ap.2017.57.0049 https://doi.org/10.1016/j.conbuildmat.2009.02.011 https://doi.org/10.1016/j.jclepro.2013.12.068 acta polytechnica ctu proceedings 26:1–4, 2020 1 introduction 2 materials and samples 3 measurement methods 4 results and discussion 5 conclusion acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0001 acta polytechnica ctu proceedings 26:1–6, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app homogenization of unreinforced old masonry wall comparison of scalar isotropic and orthotropic damage models vasco bernardo, tomáš krejčí, tomáš koudelka, michal šejnoha∗ czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: sejnom@fsv.cvut.cz abstract. masonry is a heterogeneous composite material made of bricks bonded by a mortar matrix. modeling such a material on macroscale typically calls for homogenization adopting a suitable constitutive model capable of capturing its quasi-brittle behavior. the present contribution concentrates on comparison and potential application of classical isotropic and orthotropic damage models in the framework of strain based first order numerical homogenization. as an illustrative example, a representative volume element in terms of a periodic unit cell is constructed to address the response of an unreinforced masonry (urm) structure typical of “placa” buildings (mixed masonry-reinforced concrete buildings) built in portugal. the performance of the two models is examined on the basis of macroscopic stress-strain curves constructed for both tensile and compressive loading. the selected geometrical constraints clearly identify the differences in the predictive capabilities of the two models. keywords: homogenization of unreinforced masonry wall, periodic unit cell, scalar isotropic damage model, orthotropic damage model, “placa” buildings. 1. introduction the portuguese housing stock results from more than eight centuries of history expansion. the building area consists mostly of residential buildings and it is estimated that half of the existing building stock in lisbon is represented by old masonry buildings [1]. four typologies of masonry buildings are typically recognized in the lisbon metropolitan area: buildings built before 1755, “pombalino” buildings built after the 1755 earthquake, “gaioleiro” buildings built between 1870 and 1930 and “placa” buildings 1930 to 1960 [2]. figure 1. scheme of "placa" building typology. the “placa” buildings, see fig. 1, now enjoy a particular interest as they represent 32 % of the portuguese housing stock [1]. as seen in fig. 1, the structure consists of a concrete slab simple supported by masonry walls. the exterior walls of these buildings are mainly composed of resistant stone masonry walls or by solid clay bricks with thickness greater than 0.50 m. the interior walls are usually made of solid or hollow clay bricks with thickness up to 0.30 m. the slabs are made of reinforced concrete typically with a total thickness of about 0.10 m and frequently with a single reinforcement layer and poor concrete. such building experience high vulnerability to damage when subject to seismic loads. this is attributed to insufficient strength or deformation capacity of masonry walls. it is also worth pointing out that no impact of earthquake has been considered in their design as the first portuguese seismic design regulation appeared not until 1958. predicting the bearing capacity of such buildings is thus of paramount importance, particularly in areas with a moderate to high seismic risk. a reliable prediction of the evolution of potential failure due to earthquake then strongly depends on the selected constitutive law for both concrete and masonry. herein, this issue is addressed in the light of a typically used isotropic scalar damage model [3] and more advanced orthotropic damage model allowing for an independent evolution of damage in three principal strain directions [4]. while concrete is usually assumed isotropic the masonry is only locally isotropic at the level of individual constituents whereas on macroscale it possesses an orthotropic material symmetry. to allow for treating masonry as a homogeneous anisotropic medium thus requires ho1 http://dx.doi.org/10.14311/app.2020.26.0001 http://ojs.cvut.cz/ojs/index.php/app v. bernardo, t. krejčí, t. koudelka, m. šejnoha acta polytechnica ctu proceedings mogenization. this issue is partially addressed here with reference to one specific topology of a masonry wall linked to the “placa" building. this choice is supported by available experimental data provided by the national laboratory for civil engineering in portugal (lnec) [5, 6]. the paper is organized as follows. following this introductory part, the two constitutive models are briefly reviewed in section 2. the basic grounds of the theory of homogenization are introduced next in section 3. the core of the present work appears in section 4 devoted to the numerical analysis of unreinforced masonry wall via homogenization. the principal findings are finally summarized in section 5. 2. damage models generally, damage models distinguish three states of the material virgin material, damaged material and pseudo-undamaged states. in a virgin state, the material contains no defects and behaves as linearly elastic. at points where the strength limit is exceeded the material experiences damage. in the most simple case of a scalar isotropic damage model, the damage evolution is governed by a dimensionless damage parameter ω written for one-dimensional case as ω = ad a , (1) where ad is the part of the cross section with evolved defects, i.e with the material in a damaged state and a is the total cross section area. the corresponding stress-strain relation reads σ = (1 −ω)deε, (2) where σ is the stress vector, de is the elastic stiffness matrix and ε is the vector of strain components. the scalar damage parameter ω ∈ [0; 1] characterizes the material state, where ω = 0 represents an undamaged virgin state, while ω = 1 corresponds to fully evolved defects. the transition states from 0 to 1 are described by the damage evolution law. the present formulation builds on one-dimensional traction-separation law given by [7] σ = ft exp ( − wcr uf ) , (3) where ft is the tensile strength, wcr is the crack opening and uf is the parameter controlling the slope of the softening branch. to partially avoid mesh dependency of the results typical of materials with softening the crack opening is smeared over the element as [7] κ−κe = ωκ = wcr lch , (4) where the mazars equivalent strain κ is provided by κ = √√√√ 3∑ α=1 h(εα)2, (5) where h(·) is the heaviside function, κe is the elastic part of κ and lch is the element characteristic length. combining a 1d format of eq. (2) with eqs. (4) and (3) and replacing the strain ε with the maximum equivalent strain in the loading history κ̄ yields the resulting nonlinear equation to be solved for the damage parameter ω (1 −ω)eκ̄ = ft exp ( − ωlchκ̄ uf ) , (6) for states where κ̄ exceeds the elastic threshold ε0. the scalar isotropic damage model can be used successfully for the description of 1d stress states, e.g., 1d tension or bending. a clear drawback is evident from a general multidimensional format of stress-strain relation (2) suggesting the reduction of the entire stiffness matrix even in cases when damage evolves in one direction only. the remedy can be provided by anisotropic [4] or orthotropic damage models. introducing two independent damage parameters dtα and dcα for tension and compression, respectively, gives an equivalent format of eq. (6) now written for individual principal strain directions as ( 1 −dβα ) e|εβα| = fβ exp ( − dβαlch|εβα| u β f ) , (7) where β ∈ [t,c] identifies either tensile or compressive failure. further details regarding both models can be found in [8] and [9]. 3. numerical homogenization in the present study, we limit our attention to a strain based first order homogenization only. to this end, consider a heterogeneous body loaded on its outer boundary by displacement field compatible with a macroscopically uniform strain e in an equivalent homogeneous medium. this allows us in the framework 1st order homogenization to split the local displacement and strain as [10] ∆u(x) = ∆e x + ∆u∗(x), (8) ∆ε(x) = ∆e + ∆ε∗(x), (9) where u∗(x) is the fluctuation part of the local displacement field u(x), ε∗(x) and ε(x) are the corresponding strains, and x is the spatial coordinate. next, suppose that the computational model is defined in terms of a periodic cell (puc) taking the actual brick layout into account. for “placa” building the assumed puc appears in fig. 2. the wall thickness is 250 mm, constituted by bricks with nominal dimensions of 250×120×70 mm and 10 mm for the mortar layer. because the present formulation assumes the macroscopic strain increment ∆e in eq. (8) be prescribed, the solution is searched in terms of the fluctuation rather than total displacements. the stepping stone in deriving ∆u∗ is the hill lemma written as〈 δεt(x)∆σ(x) 〉 = 0, (10) 2 vol. 26/2020 homogenization of unreinforced old masonry wall figure 2. periodic rve for the masonry wall. where the left hand side represents the volume average of virtual work done by local stress and strain fields. in the framework of finite element discretization, we approximate ∆u∗(x) in terms of their nodal values ∆r to write ∆u∗(x) = n(x)∆r, ∆ε∗(x) = b(x)∆r. (11) substituting from eq. (11)2 into (10) yields the final system of algebraic equations in the form k∆r = ∆f, (12) where k = 1 ω ∫ ω bt(x)l(x)b(x) dω, (13) ∆f = 1 ω ∫ ω bt(x)l(x)∆e dω, (14) where l is the material stiffness matrix. it is clear that in order to satisfy 〈∆ε(x)〉 = ∆e, (15) the volume average of the fluctuation part of the displacement field must disappear. this is achieved by adopting the periodic boundary conditions, which in this particular example of a rectangular puc amount to the same displacements u∗ on the opposite boundaries of puc. the rigid body motion is then constrained by simply fixing the corner displacements. note that eq. (12) can be sought as part of the solution of more complex computational framework based on multiscale analysis illustrated in fig. 3. in such a case, the actual masonry wall is replaced by an equivalent homogeneous one, so there is no need to consider all the geometrical details on macroscale. such a distinction between bricks and mortar is made on the mesoscale only being represented by a specific puc. in case of a fully uncoupled analysis, the theory of homogenization then serves to provide all the necessary data needed in the macroscopic constitutive model. the unit cell simulations then represent a virtual tester to substitute for complex or even unfeasible large scale laboratory experiments. this modeling concept is adopted henceforth providing both the effective elastic properties and various macroscopic stress-strain curves for the selected loading scenarios to compare performance of the two selected models. figure 3. framework for the homogenization of masonry wall. 4. numerical study the present section is devoted to the derivation of macroscopic response for various loading scenarios limiting attention to prescribed macroscopic uniform strains, as suggested in section 3. figure 4. masonry puc finite element model. the geometrical model displayed in fig. 2 is exploited. the corresponding, relatively fine, finite element mesh is plotted in fig. 4. this mesh consisting of four node quadrilateral elements is used first in the derivation of effective elastic properties and then in the prediction of a nonlinear response when loading the puc beyond the material elastic limits. 3 v. bernardo, t. krejčí, t. koudelka, m. šejnoha acta polytechnica ctu proceedings material e [gpa] ν [-] ft [mpa] gft [n/m] fc [mpa] ρ [kn/m3] brick 13.0 0.2 2.0 58.0 40.0 18.0 mortar 0.7 0.2 0.1 10.0 1.30 17.5 table 1. material parameters of brick and mortar. 4.1. effective elastic properties from homogenization the basic prerequisite in the application of homogenization is the knowledge of material data of individual phases as also intimated in fig. 3. in case of “placa” masonry, the computational model consists of clay bricks and mortar produced from the cement, lime and sand mixture in the proportions of 1:3. to collect the required properties, the specimens extracted during rehabilitation works from a building located in the center of lisbon were examined in the national laboratory for civil engineering in portugal (lnec), see fig. 5. the resulting material properties of both phases are listed in table 1. the same material properties for bed and head joint are assumed. further details regarding the actual experimental program are available in [5, 6, 11]. (a) (b) figure 5. axial compression tests: a) solid clay bricks, b) mortar [5]. the effective elastic plane-stress stiffness matrix is found first from four independent elasticity solutions when the unit cell is loaded in turn by each of the components of the macroscopic strain e equal to one while the remaining components vanish. the volume averages of the local stress then furnish individual columns of the homogenized effective stiffness matrix, see [12] for further details. the extracted young’s moduli and poisson’s ratios appear in table 2 clearly identifying the orthotropic material symmetry the “placa” masonry. e11 e22 g12 ν12 ν21 [gpa] [gpa] [gpa] [−] [−] 6.68 4.03 0.67 0.089 0.147 table 2. equivalent elastic properties for rve. 4.2. nonlinear response of “placa” masonry this section compares the macroscopic nonlinear response of “placa” masonry due to simple tension and compression applied independently in the x and y directions as plotted in fig. 6 also identifying the prescribed components of the macroscopic strain e. this is clearly a constrained problem as only one strain component in the principal material direction is assumed nonzero. however, because of heterogeneity all local strain and stress components become active to give rise damage evolution due to tensile failure even in case of compressive loading. (a) (b) figure 6. applied loading: a) et = {±exx, 0, 0}, b) et = {0,±eyy , 0}. the resulting macroscopic stress strain curves (σ– e) are plotted in fig. 7. the corresponding damage pattern associated with the last converged step for individual loading conditions appears in figs. 8 11. it is interesting to point out clear differences between the predictions provided by the two constitutive models observed in figs. 7(a-c). particularly the difference in the predicted strength is attributed to the fact that unlike in scalar isotropic damage model the orthotropic damage model reduces the stiffness independently in the principal strain direction according to eq. (7) providing the condition |εβα| > ε0 either for tension or compression is met. this is also supported by the evolution of stress component in the direction normal to the applied load. in this con4 vol. 26/2020 homogenization of unreinforced old masonry wall -0.002 -0.0015 -0.001 -0.0005 0 e xx [-] 0 -1.5 -1 -0.5 σ [ m p a ]σ xx orthotropic model σ yy othotropic model σ xx scalar isotropic model σ yy scalar isotropic model (a) -0.002 -0.0015 -0.001 -0.0005 0 e yy [-] 0 -1.5 -1 -0.5 σ [ m p a ] σ xx orthotropic model σ yy othotropic model σ xx scalar isotropic model σ yy scalar isotropic model (b) 0 0.0005 0.001 0.0015 0.002 e xx [-] -0.2 -0.1 0 0.1 0.2 0.3 0.4 σ [ m p a ] σ xx orthotropic model σ yy othotropic model σ xx scalar isotropic model σ yy scalar isotropic model (c) 0 0.0005 0.001 0.0015 0.002 e yy [-] 0 0.05 0.1 0.15 σ [ m p a ] σ xx orthotropic model σ yy othotropic model σ xx scalar isotropic model σ yy scalar isotropic model (d) figure 7. macroscopic stress-strain curves: a) compression (exx < 0) applied in x direction, b) compression (eyy < 0) applied in y direction, c) tension (exx > 0) applied in x direction, d) tension (eyy > 0) applied in y direction. (a) (b) (c) figure 8. damage pattern due to exx < 0 (compression): a) isotropic model, b) orthodam-11, c) orthodam-22. (a) (b) (c) figure 9. damage pattern due to eyy < 0 (compression): a) isotropic model, b) orthodam-11, c) orthodam-22. (a) (b) (c) figure 10. damage pattern due to exx > 0 (tension): a) isotropic model, b) orthodam-11, c) orthodam-22. (a) (b) (c) figure 11. damage pattern due to eyy > 0 (tension): a) isotropic model, b) orthodam-11, c) orthodam-22. 5 v. bernardo, t. krejčí, t. koudelka, m. šejnoha acta polytechnica ctu proceedings text, the results plotted in fig. 7(d) should not be surprising once realizing the geometrical arrangement of the puc. clearly, if tension is applied along the y axis, the weak mortar behaves as the weakest link. so, in this case, the stiffness along the x axis, which essentially keeps the original value in case of the orthotropic model, provides no support resulting in a similar behavior predicted by both models. although not exactly coincident with the axes of material orthotropy the damage parameter developed in the (11) principal direction (e.g., direction (11) is close to the x direction for the exx strain component prescribed) is significantly greater in comparison to the one in the second (22) principal direction, which appears negligible. it should also be mentioned that even though the two macroscopic normal stresses (σxx, σyy) figs. 7(a,b) are both compressive due to the applied constraints, e.g. exx 6= 0, while eyy = 0, the local failure is in tension because of positive strains developed as a result of material discontinuity and stress concentration at vertices of the brick and mortar interface. this causes the damage pattern in figs. 8 and 10 and figs. 9 and 11 be similar. however, because of the geometrical arrangement, the two loading conditions predict different strength limits. but as already mentioned, the maximum compressive stress reached should not be associated with the compressive strength of the masonry material. given the properties in table 1, it is not surprising that the damage localizes solely in the mortar phase for all loading scenarios. 5. conclusions the present paper was concerned with numerical modeling of masonry structure associated with the “placa” building in the framework of 1st order homogenization. apart from the derivation of effective properties of the selected masonry topology the research objectives focused on the prediction of macroscopic response employing two different constitutive models. limiting attention to tensile failure, the performance of a simple isotropic scalar damage model was compared to more advanced orthotropic damage model. the latter one allows for an independent evolution of damage in three principal directions of the local strain, thus leading in general to more stiff response as illustrated in this for both compression and tension loading condition. in this regard, the loading was introduced by prescribing the rate of macroscopic strain. only the loading in principal directions of material orthotropy was considered thus introducing some constraints to the macroscopic response. while such loading conditions appear useful in highlighting the essential differences in the behavior of the two damage models, they cannot be used in the estimation of effective parameters of the macroscopic model such as the tensile strength or the fracture energy in tension. prediction of such parameters would require combining the stress driven homogenization up to the onset of failure with a specific format of the strain-driven procedure to tract the descending part of the macroscopic stress-strain law consistent with a unidirectional stress state. this study will be the subject of our future research. acknowledgements the financial support supported by the fct project no.pd/bd/135325/2017 and by the gačr grant no. 18-24867s is gratefully acknowledged. references [1] f. tiago, j. milosevic, r. bento. seismic vulnerability assessment of a mixed masonry rc building aggregate by linear and nonlinear analyses. bulletin of earthquake engineering 14:1765–2327, 2016. [2] m. monteiro, r. bento. characterization of ”placa” buildings. icist dtc, portugal 2012. [3] g. mlani, c. valente, c. alessandri. the narthex of the church of the nativity in bethlehem: a non-linear finite eelement approach to predict the structural damage. computers and structures 207:3–18, 2017. [4] e. papa, a. taliercio. anisotropic damage model for the multiaxial static and fatigue behaviour of plain concrete. engineering fracture mechanics 1996. [5] a. i. marques, p. x. candeias, m. r. veiga, j. g. ferreira. axial compression and bending tests on old masonry walls. in 3rd international conference on protection of historical constructions. 2017. [6] a. i. marques, p. x. candeias, j. g. ferreira, m. r. veiga. caraterizacao de paredes resistentes de alvenaria antiga. in rehabend 2016, spain. 2016. [7] z. p. bažant, b. h. oh. crack band theory for fracture of concrete. matériaux et construction 16(3):155–177, 1983. [8] t. krejčí, t. koudelka, j. šejnoha, j. zeman. computer simulation of concrete structures subject to cyclic temperature loading. in b. h. v. topping, l. f. costa neves, r. c. barros (eds.), proceedings of the twelfth international conference on civil, structural and environmental engineering computing. civil-comp press, stirlingshire, united kingdom, 2009. paper 131. [9] t. koudelka, t. krejčí, j. šejnoha. analysis of a nuclear power plant containment. in b. h. v. topping, l. f. costa neves, r. c. barros (eds.), proceedings of the twelfth international conference on civil, structural and environmental engineering computing. civil-comp press, stirlingshire, united kingdom, 2009. paper 132. [10] j. c. michel, h. moulinec, p. suquet. effective properties of composite materials with periodic microstructure: a computational approach. computer methods in applied mechanics and engineering 172:109–143, 1999. [11] a. i. marques, p. x. candeias, j. g. ferreira, m. r. veiga. ensaios de compressao diagonal em paredes antigas de alvenaria de tijolo. in construcao 2018, porto. 2018. [12] m. šejnoha, j. zeman. micromechanics in practice. wit press, southampton, boston, 2013. 6 acta polytechnica ctu proceedings 26:1–6, 2020 1 introduction 2 damage models 3 numerical homogenization 4 numerical study 4.1 effective elastic properties from homogenization 4.2 nonlinear response of ``placa'' masonry 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0086 acta polytechnica ctu proceedings 26:86–93, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app euler and exponential algorithm in viscoelastic analyses of laminated glass jaroslav schmidt∗, alena zemanová czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jaroslav.schmidt@fsv.cvut.cz abstract. laminated glass combines two remarkable materials: glass and a polymer ply. while glass is stiff and brittle, the polymer ply is a rate-dependent compliant material. together, they form a material which keeps the aesthetic value of glass, and due to the polymer, no fragile collapse appears. the polymer ply exhibits timeand temperature-dependency, whereas glass suffers from brittle fracture, which makes the analysis difficult. in this article, a 2d sectional plane-stress model for the viscoelastic analysis of laminated glass is presented. this study presents one step in the development of a phase-field-based damage solver for laminated glass to select the optimal time-integration scheme for a quasistatic-analysis and later also for dynamics. the validation against experimental data is provided, and the model reduction is also discussed. keywords: laminated glass, viscoelasticity, generalised maxwell model, exponential algorithm, euler method. 1. introduction nowadays laminated glass is a popular composite multi-layer material which consists of several glass plates and polymer interlayers. the laminated process improves post-breakable behaviour compared to conventional glass but retains its aesthetic value [1]. this makes laminated glass a suitable material for transparent structures in architecture and other fields. unfortunately, despite its advantages, laminated glass is still a very brittle material which makes the simulations of the initiation and propagation of cracks difficult. this paper partially follows the project objectives specified in our research project focused on the design and advanced modelling of forced-entry and bullet resistant glass structures. one of the goals of the project is to simulate crack development in laminated glass using phase-field fracture method [2] to better understand and explain the preand post-fracture behaviour of laminated glass under impact loading. several dynamic and quasi-static experiments were performed, and the next step is the validation of the collected experimental data against numerical models for laminated glass. at this moment, the numerical solver is implemented for laminated beams and plates made of several elastic layers under quasi-static loading. this study presents another step in the development of the phase-field-based damage solver with the focus on the timeand temperature-dependent behaviour of the polymer interlayer. at this stage, the analysis deals with a quasi-static regime only. in particular, two questions concerning interlayer modelling are discussed: • what is the optimal time-integration scheme with respect to the computational time and implementation demands? • how is the global response of the model influenced by involving only certain cells of the generalised maxwell chain into the numerical analysis? moreover, this study provides also validation of material parameters of two interlayer foils obtained in the previous step of the project [3]. the paper is organised as follows: section 2 introduces the 2d plane-stress model and presents the constitutive relationships. next, in section 3, two time-integrators are recalled, and finally, section 4 provides numerical simulations, validation, and parametric study, concluded with the discussion of results in section 5. 2. laminated glass model in this section, the mechanical model for the numerical analysis is presented. the constitutive assumptions for glass and interlayer are as follows: • glass has an almost purely elastic response. therefore, it is modelled as an elastic isotropic material described by two constant: young’s modulus e and poisson’s ratio ν. • the interlayer polymer is strongly timeand temperature-dependent material effectively modelled by the generalised maxwell model [4, appendix a], see figure 1, where the material constants are the stiffnesses of springs gp and the viscosities ηp and the long-term shear modulus g∞ corresponding to the single stiffness. the constitutive relationship is expressed in the shear variables 86 https://doi.org/10.14311/app.2020.26.0086 https://ojs.cvut.cz/ojs/index.php/app vol. 26/2020 viscoelastic analysis of laminated glass instead of in the normal components as the interlayer material is almost incompressible, and shear in the interlayer is dominant in bending regime. then, τ and γ represent shear stress and strain respectively. γ(t) γe,1(t) γv,1(t) τ(t) g1 g2 g3 gp g∞ η1 η2 η3 ηp figure 1. generalised maxwell constitutive model. a practical choice for the mechanical model of laminated glass is a layer-wise model consisting of several independent beams or plates connected together at the interfaces. despite its advantages, we validate experiments against the plane-stress 2d model, where the longitudinal cross-section is discretizted. one of the reasons for this approach is to check if the throughthe-thickness compression of the interlayer is really negligible and can be neglected. 3. time integrators the brief overview of the exponential and backwardeuler integrators is presented in this section. 3.1. governing equations the following model assumptions are considered: • a special two-dimensional state of stress called plane stress is assumed to reduce the originally three-dimensional task and subsequently the computational cost. • because all of the laminated glass samples were simply-supported, the analysis assumes only small deformations and deflections. the comparison with experiments shows that this assumption is sufficient as no stabilisation due to membrane forces occurred. • we assume that the poisson ratio of the polymer interlayer is constant. for further comparison with a constant bulk modulus assumption see also [5]. as mentioned before, the polymer material point is described by the generalised maxwell chain which consists of one single linear spring with stiffness g∞ and several maxwell cells. each cell is characterised by a spring with stiffness gp and a damper with viscosity ηp. the cells are indexed by p ∈ p, where p = 〈0,p〉 ⊂ n. the time integrators are used only for the viscoelastic interlayer and glass is considered as rate-independent. under these assumptions, the behaviour of the interlayer can be described by a weak form of the equilibrium equation ∫ ωfoil δε :  g∞dν : ε + ∑ p∈p σp   dω − δfext = 0, (1) complemented by ordinary differential equations (ode) σ̇p gp + σp ηp = dν : ε̇, p ∈p. (2) in the former equations δfext stands for the virtual work done by the external loads on the virtual displacement δu, and the small strain tensor ε is obtained as the symmetric displacement gradient ε = ∇su. the dimensionless tensor dν corresponds to the stiffness tensor of an isotropic material expressed by a unit shear modulus and the poisson ratio of ν. the stress carried by p-th cell is denoted by σp. on the other hand, the governing equation for the deformation of glass does not contain viscous stress tensors σp, and the corresponding weak form becomes∫ ωglass δε : gglassdν : ε dω − δfext = 0. (3) the implementation for laminated glass, which combines the elastic glass domain and the viscoelastic interlayer, is described in section 4. equations (1)–(2) contain both space and time gradients. the space discretization is captured by using conventional finite element methods (fem), whereas the type of the time discretization is a part of this analysis. two discretization options are recalled in the following subsection. 3.2. numerical integrators the common feature for all presented time integrators is the discretization into a set of time instants t = {0, . . . , tmax}, where tmax is the largest time of interest. backward euler method for the backward euler method, we approximate the time gradient by the firstorder backward finite difference •̇(ti+1) ≈ •(ti+1) −•(ti) ∆ti . (4) to shorten the notation, the time increment ∆ti is defined as ti+1 − ti, and the explicit time dependency is omitted and replaced by a lower index, therefore •(ti) ≡ •i. if we use the above-mentioned gradient approximation of the ode evaluated in time instant ti+1, we get σp,i+1 −σp,i gp∆ti + 1 ηp σp,i+1 = 1∆ti dν : (εi+1 −εi) . (5) then, we can express explicitly σi+1 as σp,i+1 = 1 ζp,i (dν : (εi+1 −εi)gp + σp,i) , (6) 87 jaroslav schmidt, alena zemanová acta polytechnica ctu proceedings where ζp,i = 1 + ∆ti/τp is an auxiliary dimensionless parameter. the stress tensor can be directly substituted into the equilibrium equation to get the weak form for εi+1 and the resulting equation ∫ ωfoil δε :  g∞ + ∑ p∈p gp ζp,i   dν : εi+1 dω − ∫ ωfoil δε : ∑ p∈p gp ζp,i dν : εi dω (7) + ∫ ωfoil δε : ∑ p∈p σp,i ζp,i dω −δfext = 0 governs the fe analysis for εi+1 determination. then, the stress tensors are updated using (6). this procedure is repeated for all time instants. exponential algorithm the main difference for the exponential algorithm compared with the euler method is the discretization of the stress tensor. as we suppose that the strain is distributed linearly within the time interval 〈ti, ti+1〉, the viscous stress tensor is solved exactly using the corresponding ode. the linearly distributed strain induces a constant strain rate, and the governing equation for the maxwell unit becomes σ̇p(ti + si) gp + σp(ti + si) ηp = 1 ∆ti dν : (εi+1 −εi) , (8) where si is an auxiliary time variable starting from ti and going forward in time. the relation between the original time t and the variable si is si = t− ti. considering the boundary condition σp(si = 0) = σp,i, the stress evaluated in t = ti+1 reads as σp,i+1 = σp,ie −∆ti τp + ηp ∆ti ( 1 − e− ∆ti τp ) dν : (εi+1 −εi). (9) the backward substitution into the equilibrium condition (1) in ti+1 gives ∫ ωfoil δε :  g∞ + ∑ p∈p ηp ∆ti ξp,i   dν : εi+1 dω − ∫ ωfoil δε : ∑ p∈p ηp ∆ti ξp,idν : εi dω (10) + ∫ ωfoil δε : ∑ p∈p e− ∆ti τp σp,i dω −δfext = 0, where another auxiliary variable ξp,i = 1−e −∆ti τp is introduced. 4. numerical results 4.1. example all calculations were performed on a longitudinal crosssection of a laminated glass beam assuming a planestress state, see figure 2. the dimensions of the beam are: the thicknesses h1 = h3 = 10 mm, h2 = 0.76 mm, the overhanging edges l1 = l3 = 30.0 mm and their span l2 = 1040.0 mm. h1 h2 h3 l1 l2 l3 figure 2. scheme of laminated glass sample. the laminated glass samples were simply-supported and loaded by prescribed uniformed pressure in a vacuum chamber. the time evolution of pressure was almost the same for the two tested samples with eva and pvb interlayer, see figure 3. at the beginning of loading, the pressure increase is almost quadratic, whereas later the slope becomes linear. 0 5 10 15 20 25 30 35 40 time [s] 5 0 5 10 15 20 25 30 35 40 pr es su re lo ad [k pa ] pressure load figure 3. time evolution of pressure recorded during the experimental testing. 4.2. material parameters glass is treated as a perfectly elastic material described by the young modulus eglass = 76.6 gpa obtained from indentation tests and by the poisson ratio νglass = 0.22. whereas the interlayer polymer is effectively described by the generalised maxwell model to capture its viscoelastic time/temperature dependent nature. in the tested samples, an ethylene-vinyl acetate (eva) and a polyvinyl butyral (pvb) foils were used as the interlayer material, with the material constants taken from [3] and given in table 1. the poisson ratio was assumed to be constant, νfoil = 0.49. 88 vol. 26/2020 viscoelastic analysis of laminated glass eva pvb τp [s] gp [kpa] gp [kpa] 10−9 6933.9 – 10−8 3898.6 – 10−7 2289.2 – 10−6 1672.7 – 10−5 761.6 1,782,124.2 10−4 2401.0 519,208.7 10−3 65.2 546,176.8 10−2 248.0 216,893.2 10−1 575.6 13,618.3 100 56.3 4988.3 101 188.6 1663.8 102 445.1 587.2 103 300.1 258.0 104 401.6 63.8 105 348.1 168.4 106 111.6 – 107 127.2 – 108 137.8 – 109 50.5 – 1010 322.9 – 1011 100.0 – 1012 199.9 – table 1. prony series for generalised maxwell model complemented with the long-term moduli g∞ = 682.18 kpa for evalam 80-120 (eva-based foil) and g∞ = 232.26 kpa for trosifol bg r20 (pvbbased foil) for t0 = 20 ◦c from [3]. the room temperature during the experimental testing was 24◦c. for the polymer ply, the temperature dependency is incorporated into the material model using the time-temperature superposition principle [6, chapter 11]. time and temperature are tied together via a multiplicative parameter at as τp(t) = at(t) ηp gp . (11) the parameter at can be evaluated through the william–landel–ferry equation (wlf), [7], log10 at = − c1(t −t0) c2 + t −t0 , (12) where the constants c1 and c2 are additional material parameters, and t0 is a reference temperature. for the reference temperature t0 = 20 ◦c, the calibration process described in [3] provides the following values: c1 = 339.1 and c2 = 1185.8◦c, or c1 = 8.635 and c2 = 339.102, for eva or pvb foil respectively. 4.3. solver for a given time instant ti+1, the unknown displacement field u(ti+1) is calculated numerically by the conventional finite element method. the part of the discretized domain is illustrated in figure 4, where the red colour represents glass layers and the blue colour the interlayer. due to the disproportionality between the glass and interlayer thicknesses, the mesh is refined near the centerline with at least two elements through the ply thickness. moreover, the quadratic basis functions were employed. the computational fem library fenics was used to quickly implement the presented models for all numerical experiments. this solver with a high-level python interface provides an efficient tool for solving partial differential equations [8]. figure 4. finite element mesh: glass domain in red and the interlayer in blue. for modelling of the glass/polymer coupling, an indicator function i(x) was introduced as follows i(x) = { 0, if x ∈ ωglass, 1, if x ∈ ωfoil. (13) all viscous members of the governing equations, section 3.1, are multiplied by this function. therefore, the rate-dependent expressions appear only in the interlayer domain ωfoil. let us highlight that no delamination, glass fracture, or damage of polymer is considered in this study. 4.4. validation and comparison the main advantage of the exponential algorithm is that the time interval can be covered by a series of logarithmically increasing steps without significant loss of accuracy. therefore, the computational cost and the number of time increments can be significantly reduced, for example for long-term stress relaxation or creep quasi-static tests. the results for 20 logarithmically or uniformly distributed time steps are presented in figure 5. the reference solution correspondes to the backward euler method with the time step ∆t = 0.05 s. the detailed comparison of responses on the right side shows the advantage of the exponential algorithm for logarithmically increasing time steps, as was supposed. on the other hand, the performance of both methods is the same for uniformly distributed time steps. for uniformly distributed time steps, three different time increments were tested, i.e., 0.05 s, 0.5 s and 2.0 s. the numerical responses are presented together with the experimental data in figure 6. this validation shows that even the numerical response with 89 jaroslav schmidt, alena zemanová acta polytechnica ctu proceedings 0.0 0.5 1.0 1.5 2.0 2.5 3.0 time [s] 0.00000 0.00005 0.00010 0.00015 0.00020 0.00025 0.00030 0.00035 0.00040 0.00045 m ax im um d ef le ct io n [m m ] reference exp. log be log 0.700 0.705 0.710 0.715 0.720 time [s] 0 2 4 6 8 10 m ax im um d ef le ct io n [m m ] 1e 7+3.4e 5 detail (a) logarithmically distributed time steps 0.0 0.5 1.0 1.5 2.0 2.5 time [s] 0.00000 0.00005 0.00010 0.00015 0.00020 0.00025 0.00030 0.00035 0.00040 0.00045 m ax im um d ef le ct io n [m m ] reference exp. be 0.7880 0.7885 0.7890 0.7895 0.7900 time [s] 2.0 2.5 3.0 3.5 4.0 4.5 5.0 5.5 6.0 m ax im um d ef le ct io n [m m ] 1e 7+4.2e 5 detail (b) uniformly distributed time steps with ∆t = 0.125 s figure 5. time evolution of deflection of laminated glass sample under uniformly distributed pressure with quadratically increasing intensity according figure 3 obtained by: reference – backward euler method with ∆t = 0.05 s, exp – exponential algorithm, and be – backward euler method. 15 20 25 30 time [s] 0 5 10 15 m ax im um d ef le ct io n [m m ] exponential backward euler experiment ∆t = 0.05 s 15 20 25 30 time [s] 0 5 10 15 exponential backward euler experiment ∆t = 0.5 s 15 20 25 30 time [s] 0 5 10 15 exponential backward euler experiment ∆t = 2.0 s (a) response for samples with eva foil 14 16 18 20 22 24 26 28 time [s] 0 5 10 15 m ax im um d ef le ct io n [m m ] exponential backward euler experiment ∆t = 0.05 s 14 16 18 20 22 24 26 28 time [s] 0 5 10 15 exponential backward euler experiment ∆t = 0.5 s 14 16 18 20 22 24 26 28 time [s] 0 5 10 15 exponential backward euler experiment ∆t = 2.0 s (b) response for samples with pvb foil figure 6. validation of numerical responses provided by the exponential and backward euler method against the experimental measurement for three different time steps: ∆t = 0.05 s, ∆t = 0.5 s and ∆t = 2.0 s. 90 vol. 26/2020 viscoelastic analysis of laminated glass 15 20 25 30 time [s] 0 5 10 15 m ax im um d ef le ct io n [m m ] dt=0.05s dt=0.5s dt=2.0s dt=5.0s dt=10.0s b-euler figure 7. numerical response of backward euler method for different time steps. the coarser time steps is in good agreement with the measured deflections. the dashed line representing the exponential algorithm almost coincides with the solid one for the backward euler method in all cases. for larger time steps, the differences of numerical predictions against the measured experimental response would increase, see figure 7. on the other hand, we aim to describe the crack initiation and propagation in glass layers, which requires small time steps. it results from this comparison that both methods are able to capture the time/temperature-dependent response of polymer interlayers well. further, the validation against experimental data proved that the parameters of the generalised maxwell model for both foils, obtained from material tests and summarised in [3], provides very good predictions for quasi-static behaviour of laminated glass samples with eva-based and pvb-based foils of the same type. the response is just slightly stiffer at the beginning of loading for the sample with eva foil and a slightly softer before the fracture of glass for the sample with pvb. however, the largest error is under 5%. 4.5. sufficient number of parameters the comparison of the viscoelastic response of laminated glass samples with three elastic simulations in figure 8 demonstrates the influence of shear coupling of glass plies due to the interlayer. the behaviour of the polymer was approximated as an elastic one with three different values of shear moduli. the response of the laminated glass sample is often bounded by the behaviour of a monolithic glass plate with the same overall thickness and by the response of two independent plies representing the two glass plates without any coupling. we narrowed this range, assuming a constant long-term or initial shear modulus, g∞ or g0. for the sample with eva foil, a reasonable approximation of the viscoelastic response was found using the shear modulus of the interlayer corresponding to a half of the loading time g(t/2). for pvb, this assumption gives still a softer prediction of deflections compared with the measured data. in the next stage, we tried to reduce the number of spring-dashpot cells as each of these maxwell cells is represented by a second-order viscous stress tensor. therefore, it is desirable to balance the number of cells and save the computational time. the exponential and backward euler methods provided the same response for the fine time step. thus, only the backward euler method was used for the analysis. the influence of the individual maxwell cell can be assessed using the multiplicator ζp,i. it can be seen from equation (7) that the significant viscous stresses are those where the expression 1/ζp,i is not close to zero. recalling the definition of ζp,i, 1 ζp,i = 1 1 + ∆ti τp = τp τp + ∆ti , (14) the whole fraction is close to one if the relaxation time is much more greater than the given time step τp � ∆ti, and the viscous stress is activated. on the other hand if τp � ∆ti, the denominator is large and the whole fraction is close to zero. to verify this observation we perform two numerical tests. figure 9 shows that if we remove the low-relaxation-time cells, the response is not affected (left), and conversely, removing of cells with large relaxation time have significant influence on the response (right). the reference solution plotted by the red line corresponds to the original generalised maxwell model with all cells, table 1. the number of cells was gradually decreased. figure 9 (left) illustrates the case, for which the cells with the smallest relaxation times were neglected. it can be seen that neglecting even first seven cells for eva or three for pvb with the smallest relaxation times did not affect the solution as the time evolution of deflection is indistinguishable from the response of the full generalised maxwell model. on the other hand, removing a few maxwell cells with the largest relaxation times significantly affected the deflection of the sample, figure 9 (right). this is caused by the fact that we start in the proposed formulation with the same long-term modulus. 5. conclusions the numerical analysis of laminated glass samples under quasi-static loading is presented in this paper and validated against experimental measurements. the laminated glass plate is discretized as a 2d longitudinal cross-section with two elastic glass layers and the viscoelastic interlayer. the polymer ply was characterised by the generalised maxwell model with parameters derived from material tests in [3]. the conclusions from this study are: • both tested methods for time integration, i.e., the exponential algorithm and the backward euler method, can be used for the subsequent finite element analysis of glass fracture. the difference in results is negligible for values of time steps needed in proposed phase-field formulation. 91 jaroslav schmidt, alena zemanová acta polytechnica ctu proceedings 15 20 25 30 time [s] 0 5 10 15 m ax im um d ef le ct io n [m m ] linear g_inf linear g_0 linear g(t/2) exponential ∆t = 0.05 s 15 20 25 30 time [s] 0 5 10 15 m ax im um d ef le ct io n [m m ] linear g_inf linear g_0 linear g(t/2) exponential ∆t = 0.05 s (a) sample with eva foil (b) sample with pvb foil figure 8. comparison of the viscoelastic response of laminated glass samples obtained by the exponential algorithm with the elastics ones with three different shear moduli of interlayer assumed. 15 20 25 30 time [s] 0 5 10 15 m ax im um d ef le ct io n [m m ] n=0 n=5 n=10 n=15 n=22 (ref.) 15 20 25 30 time [s] 0 5 10 15 m ax im um d ef le ct io n [m m ] n=0 n=5 n=10 n=15 n=22 (ref.) (a) samples with eva foil 15 20 25 30 time [s] 0 5 10 15 m ax im um d ef le ct io n [m m ] n=0 n=4 n=8 n=11 (ref.) 15 20 25 30 time [s] 0 5 10 15 m ax im um d ef le ct io n [m m ] n=0 n=4 n=8 n=11 (ref.) (b) samples with pvb foil figure 9. time evolution of deflections predicted by the backward euler method with time step ∆t = 0.5 s: reference – numerical simulation with all maxwell (spring-dashpot) cells or reduced – with n maxwell cells if the cells with the smallest relaxation times (left) or the largest relaxation times (right) are removed. 92 vol. 26/2020 viscoelastic analysis of laminated glass • for laminated glass samples with eva-based or pvb-based foils, the formerly fitted material parameters provides reliable numerical predictions of deflections. the errors in deflections were under 5%. • the computational cost can be reduced without the loss of accuracy using a constant shear modulus for the interlayer g(t/2) for eva-based samples. for pvb-based samples, three maxwell cells with the smallest relaxation times can be neglected without an observable change in response of the sample. finally, let us conclude that this study was focused on the comparison of the response of laminated glass before fracture with emphasis on the time/temperature-dependent behaviour of polymer interlayers. no phase-field description of glass fracture is presented and discussed in this paper, and also no damage of polymer foils is assumed. the reason for this is the fact that based on the phase-field model formulation and setting, the linear elastic phase can also be affected by the damage evolution if the growth of damage starts immediately when the strain becomes nonzero. then, the model does not reproduce the linear elastic behaviour of laminated glass in the first stage, and the stress-strain diagram becomes curved. therefore, the phase-field model for brittle fracture of glass will be discussed separately to avoid the possible deviation from linearity and misinterpretation of results. acknowledgements this publication was supported by the czech science foundation, the grant no. 19-15326s and by the grant agency of the czech technical university in prague, grant no. sgs19/033/ohk1/1t/11. references [1] m. haldimann, a. luible, m. overend, et al. structural use of glass. structural engineering documents. international association for bridge and structural engineering, 2008. [2] b. bourdin, g. a. francfort, j. j. marigo. the variational approach to fracture. journal of elasticity 91:5—-148, 2008. doi:10.1007/s10659-007-9107-3. [3] t. hána, t. janda, j. schmidt, et al. experimental and numerical study of viscoelastic properties of polymeric interlayers used for laminated glass : determination of material parameters. materials 12(14), 2019. doi:10.3390/ma12142241. [4] z. p. bažant, m. jirásek. creep and hygrothermal effects in concrete structures. springer, 2018. doi:10.1007/978-94-024-1138-6. [5] a. zemanová, j. zeman, m. šejnoha. comparison of viscoelastic finite element models for laminated glass beams. international journal of mechanical sciences 131-132:380–395, 2017. doi:10.1016/j.ijmecsci.2017.05.035. [6] j. d. ferry. viscoelastic properties of polymers. john wiley and sons, 1980. doi:10.1149/1.2428174. [7] m. l. williams, r. f. landel, j. d. ferry. the temperature dependence of relaxation mechanisms in amorphous polymers and other glass-forming liquids. journal of the american chemical society 77(14):3701–3707, 1955. doi:10.1021/ja01619a008. [8] m. s. alnæs, j. blechta, j. hake, et al. the fenics project version 1.5. archive of numerical software 3(100), 2015. doi:10.11588/ans.2015.100.20553. 93 https://doi.org/10.1007/s10659-007-9107-3 https://doi.org/10.3390/ma12142241 https://doi.org/10.1007/978-94-024-1138-6 https://doi.org/10.1016/j.ijmecsci.2017.05.035 https://doi.org/10.1149/1.2428174 https://doi.org/10.1021/ja01619a008 https://doi.org/10.11588/ans.2015.100.20553 acta polytechnica ctu proceedings 26:86–93, 2020 1 introduction 2 laminated glass model 3 time integrators 3.1 governing equations 3.2 numerical integrators 4 numerical results 4.1 example 4.2 material parameters 4.3 solver 4.4 validation and comparison 4.5 sufficient number of parameters 5 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0245 acta polytechnica ctu proceedings 33:245–249, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague durability of high belite cement as new technical solution for concrete sara iricoa, ∗, sabine mutkea, federica bertolab, daniela gastaldib, livio capellib, fulvio canonicoa, b a wilhelm dyckerhoff institut for building materials technology, dyckerhoff gmbh, dyckerhoffstraße 7, 65203 wiesbaden, germany b buzzi unicem innovation lab and technology (built), via restano 3, 13100, vercelli, italy ∗ corresponding author: sara.irico@dyckerhoff.com abstract. the study presents an investigation on the durability performance of concretes produced with an industrial scale high belite cement. this binder, thanks to the high amount of c-s-h gel and the lower c3s/c2s ratio, is characterized by high durability properties. good resistance to freeze/thaw cycles, very low chloride migration and excellent resistance to sulfate attack were demonstrated. the accelerated carbonation test revealed that the belite cement, due to the high c2s phase amount, is able to increase its strength over time when exposed to high co2 concentration. durability and strength results were compared with those of cem i, cem ii, and cem iii. in addition, a comparison of the performance of a self-compacting concrete produced with high belite cement and cem i (as reference) is also presented. the results show that the self-compacting concrete with the high belite cement is able to reach sufficient strengths already after 1 day of hydration, with a low heat development and improved durability properties, compared to cem i. keywords: belitic cement, low heat cement, low heat concrete. 1. introduction new technical solutions aimed at building sustainable and durable structures need a synergy between the research on innovative binders, with reduced co2 emissions, and the development of the advanced concrete technology. today, the use of supplementary cementitious materials (scms) is the main strategy to lower the environmental impact of the cement and concrete production. however, the long-term availability of fly-ash and slag will drop in the next future and has to be considered. while, limestone is the most used scm being economic and widely available, but its use in concrete is an issue concerning durability requirements [1]. in this study, a high belite cement (hbc) is proposed as technical solution for concretes requiring low heat development and long service life. this type of binder is characterized by a reduction of co2 footprint around 8-10% when compared to conventional portland cement, thanks toa reduced limestone demand as raw material and a lower burning temperature needed for clinker production.. the larger amount of c-s-h gel developed by the belite phase than alite is the basis to expect an improvement in durability performance, a possible driving force to move from portland clinker to belite clinker or future blends with scms. the hbc (4540 cm2/g) used in this study was produced at industrial scale by buzzi unicem s.p.a. the c2s content is higher than 54 wt.-% and the heat of hydration is lower than 227 j/g during 7 days. performance and durability of the hbc were compared with those of cem i 52.5n (3600 cm2/g), , cem ii/b-ll 32.5r (5690 cm2/g), and cem iii/a 52.5n lh (4150 cm2/g), as references. 2. results 2.1. compressive strengths related to the heat of hydration an undoubted advantage in the use of hbc is represented by its very low heat of hydration; in fact, as shown in figure 1 it exhibits the lowest value between the different cements, also considerably lower than cem iii. this fact makes hbc highly interesting for mass concrete applications and for possible blends with portland clinker and scms when low heat cement is required. figure 1 shows also the comparison of the cements in terms of compressive strength, mortars were produced according to the en 196-1 standard. after 1 day hbc develops a slightly lower compressive strength than cem ii, due to the slower hydration rate of the c2s phase,and responsible for the low heat development. at 7 days the compressive strength of hbc and cem ii are almost comparable, while from 28 days on hbc exhibits higher results. 2.2. high belite cement for concrete at exposure classes xd, xs and xf the concretes were produced according to the boundary conditions dictated in the german institute for 245 https://doi.org/10.14311/app.2022.33.0245 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en s. irico, s. mutke, f. bertola et al. acta polytechnica ctu proceedings figure 1. comparison of the mortar compressive strength development correlated with the heat of hydration referred to the mass of cement. figure 2. compressive strength of the c1 concretes after 28 days of curing. construction technology (dibt) approval process for cements. two concrete compositions were considered in this study. the resistance to chloride penetration, according to the baw code of practice (similar to nt built 192), was performed on concrete c1 produced by using rhine gravel and sand with a16/b16 grading curve, 320 kg/m3 of cement, and a w/c ratio of 0.5. no plasticizers were used in order to avoid their influence. figure 2 shows the compressive strengths of concrete c1 produced with the different cements used at 28 days of curing; hbc shows exhibits a higher strength compared to cem ii, but lower than the other two cements. the resistance to the freeze/thaw cycles immersed in 3% sodium chloride solution, according to the cdf method (din en 12390-9), was instead performed on concrete c2 produced with: a16/b16 grading curve, 320 kg/m3 of cement , a w/c ratio of 0.5, and an air figure 3. chloride migration coefficients of c1 concretes. the solid square symbols show the compressive strength values at 28 days of hydration of mortars. void content of 4.5 ± 0.5 vol%. as shown in figure 3, hbc shows a higher resistance against chloride migration compared to cem i and cem ii, but lower than cem iii. the migration coefficient of hbc is low enough for hydraulic engineering applications (≤ 10 × 10−12 m2/s for xs1-2, xd1-2 or ≤ 5 × 10−12 m2/s for xs3 and xd3, in accordance with [2]). few information are reported in literature about the resistance of hbc to freeze/thaw cycles. the results, shown in the graph in figure 4, reveal that hbc behaviour is comparable to that of cem i; both cements are suitable for concrete applications with xf4 exposure class, presenting a scaling after 28 cycles lower than the threshold limit. on the contrary, both cem ii and cem iii show a weak resistance against 246 vol. 33/2022 durability of high belite cement figure 4. resistance to freeze/thaw attack with deicing salt (cdf method). figure 5. expansion due to sulfate attack at 20 ◦c and 5 ◦c. the stress induced by the 28 freeze/thaw cycles going beyond the threshold limit. these data confirm that the freeze/thaw resistance is a "weakness" property of limestone and slag cements, as well known. the resistance to freeze/thaw cycles is often an obstacle to the implementation of sustainable binders with lower clinker factor. 2.3. sulfate resistance the sulfate resistance of the cements was tested according to the sva flat prism (4 × 1 × 16 cm) method. the expansion of the prisms cured in a saturated ca(oh)2 solution was used as reference in order to evaluate the influence of sulfate attack at 20 ◦c and 5 ◦c. hbc has an excellent resistance against the aggression of sulfate, as confirmed by sui et al. [3], and by the experimental results in figure 5 where hbc exhibits the lowest expansion compared to the other cements. the resistance of portland cement exposed to sulfate solution is related, in general, to the c3a content, by the way a high amount of portlandite development can make concrete more vulnerable to sulfate attack, with the consequence formation of gypsum. the positive influence of the low c3s/c2s ratio (figure 5), often not considered in the interpretation of the results, is a significant factor in the choice of a sulfate resistant cement [4]. the very high expansion of the cem ii at 5 ◦c is probably due to the formation of thaumasite [5]. 2.4. accelerated carbonation in order to compare the depth of carbonation in the different cements considered, mortar samples were prepared in accordance with en 196-1, stored in water up to 7 and 28 days, and then exposed to accelerated carbonation (chamber with 4% of co2 at 20 ◦c, and 70% rh) for 28, 56, 90, 140 days. the carbonation depth was determined at each due date with 247 s. irico, s. mutke, f. bertola et al. acta polytechnica ctu proceedings figure 6. accelerated carbonation (4% co2 at 20 ◦c, and 70% rh) and influence on the strength development. the phenolphthalein test in accordance with rilem recommendation. in addition, prior to determine the co2 penetration, on the same samples, compressive strengths were also evaluated to understand the effect of the belite content and the influence of carbonation curing on the physicochemical properties (figure 6). the carbonation-cured hbc has a significant growth in the compressive strength during the exposure time. jang at al. [6] demonstrated that the co2 uptake capacity is proportional to the belite content and most of the belite phase is consumed by the carbonation reactions during the curing, resulting in the production of calcite. they observed that the carbonation curing of cement with high belite content leads to a more dense microstructure (with lower connected porosity) than those with high alite content. the hypothesis states that the carbonation reactions occurring at the inner layer of an unreacted belite particle are likely more rapid than that of the hydration reaction, resulting in the densification of the microstructure between the inner layer of c-s-h and the outer layer of unreacted belite particles, which significantly influences the strengths. the depth of carbonation of hbc is proved to be strongly dependent on its grade of hydration, in fact when cured just for 7 days it is considerably higher than when it is cured for 28 days. 2.5. high performance concrete with hbc a challenge in the application of hbc is represented by the slower strength development compared to conventional cements. research on cement science is addressing efforts in order to overcome this barrier to the implementation. on the other hand, it must be said that in case of new eco-binders or innovative cements, not included in the standard en 197-1, the evaluation of the performance considering standard mortars with a w/c of 0.5 (en 196-1) could be not sufficient to explore their potential. on this purpose, progresses in concrete technology play a relevant role in order to match the needs of implementing ecobinders characterized by lower environmental impact with requirements on compressive strength, rheology and durability. the behavior of a self-compacting concrete (scc) with 420 kg/m3 of hbc, 100 kg/m3 of fly-ash, and a w/c ratio of 0.36 was compared with a scc made with cem i 52.5r, as reference. the compressive strength of the hbc-scc after 1 day reaches 20 mpa, and after 7 days is comparable to that of cem i-scc. hbc has also the great advantage of a very low heat of hydration. the durability of the two sccs was evaluated in terms of freeze/thaw resistance and chloride penetration. the scaling due 248 vol. 33/2022 durability of high belite cement to freeze/thaw cycles (according to the cdf test) is in both cases very low: 450 g/m2 for cem i-scc, and just 192 g/m2 for hbc-scc. the chloride migration of hbc-scc is lower than that of cem i-scc, 2.2 × 10−12 m2/s and 4.7 × 10−12 m2/s, respectively. 3. conclusions the durability of concretes made with an industrially produced hbc was explored. the high resistance against chloride migration and freeze/thaw scaling of hbc, in comparison with cem i, cem ii, and cem iii (as reference), were evaluated on concretes produced according to the boundary conditions dictated in the german institute for construction technology (dibt) approval process for cements. hbc revealed to be suitable for exposure classes xs3, xd3 and xf4, thanks to its good resistance against freeze/thaw cycles, and its extraordinary ability to counteract the chloride penetration. almost no expansion was measured on hbc samples due to sulfate attack. the accelerated carbonation showed that hbc, thanks to the high amount of belite phase (higher than 54wt%), increases its compressive strengths during the co2 exposure. this interesting phenomenon needs to be deeply investigated at mineralogical and microstructural level, as well as the high durability properties of this cement related to the c-s-h gel formation. the durability of hbc is probably the main driving force for the future applications of this binder, together with its reduced environmental impact, and its low of heat of hydration. the study also demonstrated that when formulated as self-compacting concrete, with low w/c, hbc can reach already after 1 day enough compressive strength (higher than 20 mpa) to work efficiently. the potentialities of this binder need to be further explored also in blends with scms in order to develop sustainable and more durable future cementitious binders. references [1] w. neufert, i. reuken, g. weber g, et al. erforschung des dreistoffgemisches klinker, hüttensand und kalksteinmehl mit dem ziel der absenkung des klinkeranteils im zement zur verminderung der co2 emissionen in der zementproduktion, abschlussbericht reduzierung der umweltwirkung der betonbauweise, 2016. https://www.dbu.de/opac/ab/dbu-abschlussbe richt-az-28790-02.pdf. [2] bundesanstalt für wasserbau (hg.). bawmerkblatt chlorideindringwiderstand von beton (mcl), karlsruhe: bundesanstalt für wasserbau (baw-merkblätter, -empfehlungen und -richtlinien), 2012. https://henry.baw.de/bitstream/handle/20. 500.11970/102465/bawmerkblatt_chlorideindringwi derstand_mcl_2012_k.pdf?sequence=1&isallowed=y. [3] s. tongbo, f. lei, w. zhaijun, et al. properties of belite-rich portland cement and concrete in china. journal of civil engineering and architecture 9(4):384-392, 2015. https://doi.org/10.17265/1934-7359/2015.04.002. [4] a. neville. the confused world of sulfate attack on concrete. cement and concrete research 34(8):1275-96, 2004. https: //doi.org/10.1016/j.cemconres.2004.04.004. [5] m. collepardi. a state-of-the-art review on delayed ettringite attack on concrete. cement and concrete composites 25(4-5):401-7, 2003. https://doi.org/10.1016/s0958-9465(02)00080-x. [6] j. g. jang, h. k. lee. microstructural densification and co2 uptake promoted by the carbonation curing of belite-rich portland cement. cement and concrete research 82:50-7, 2016. https: //doi.org/10.1016/j.cemconres.2016.01.001. 249 https://www.dbu.de/opac/ab/dbu-abschlussbericht-az-28790-02.pdf https://henry.baw.de/bitstream/handle/20.500.11970/102465/bawmerkblatt_chlorideindringwiderstand_mcl_2012_k.pdf?sequence=1&isallowed=y https://doi.org/10.17265/1934-7359/2015.04.002 https://doi.org/10.1016/j.cemconres.2004.04.004 https://doi.org/10.1016/s0958-9465(02)00080-x https://doi.org/10.1016/j.cemconres.2016.01.001 acta polytechnica doi:10.14311/app.2020.27.0112 acta polytechnica 27(0):112–115, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app local nano-mechanical properties of cross-linked polybutylene martin ovsík∗, michal staněk, adam dočkal, petr fluxa tomas bata university in zlin, department of production engineering, nám. t. g. masaryka 5555, 760 01 zlín, czech republic ∗ corresponding author: ovsik@utb.cz abstract. cross-linking is a process in which polymer chains are associated through chemical bonds. the cross-linking level can be adjusted by the irradiation dosage and often by means of a cross-linking booster. the polymer additional cross-linking influences the surface nano and micro layers in the way comparable to metals during the thermal and chemical-thermal treatments. polybutylene terephthalate (pbt) can be found in a group of structural polymers, which are often used in industry, especially in automotive. applying the technology of electron radiation induces a creation of 3d network structure, which improves the local mechanical properties. these were later measured by a depth sensing indentation (dsi) test. this state of the art method is based on immediate detection of indentation depth in relation to applied force. the creation of 3d network caused an increase in nano-mechanical properties values, such as indentation hardness and indentation modulus, in comparison to the virgin material. the indentation hardness rose by 80%, while the indentation modulus elevated by 62%. the selected structural materials, e.g. pbt, were modified by the electron irradiation in a positive way and as such could be moved to a group of high performance materials. keywords: cross-linked, electron radiation, hardness, nano-indentation, polybutylene terephthalate. 1. introduction polybutylene terephthalate (pbt) is a semicrystalline polymer. some of the remarkable properties of pbt are as follows [1–3]: • good chemical resistance • high dielectric strength, excellent electrical properties • high heat resistance and temperature performance • good strength and modulus at elevated temperatures • very good processability • flame resistance and ease of automated soldering. the major application of pbt is in the automotive sector. it is used for the manufacturing of electronic components, parts of instrument panels, and housings of automotives. exterior components, namely bumper fascias, mudguards, door handles, mirror housings, and wiper arms are also produced from pbt. its electrical applications include connectors, smart network interface devices, power plugs and electrical components, switches and controls, circuit breaker enclosures, and outdoor telecommunication enclosures. it has also found application in fiber-optic tubing and electrical appliances where the components require a high surface gloss [4–7]. e-beam crosslinking is a powerful tool used to improve the properties of a wide range of polymers in the creation of value-added specialty products. the crosslinking of polymers through electron-beam processing changes a thermoplastic material into a thermoset. when polymers are crosslinked, the molecular movement is severely impeded, making the polymer stable against heat. crosslinking is the interconnection of adjacent long molecules with networks of bonds induced by chemical treatment or electron-beam treatment [8–11]. this study is concerned with the effect of ionizing radiation upon the local nano-mechanical properties of filled pbt. the main goal was to evaluate the influence of cross-linking (creation of the 3d network) on the local nano-mechanical properties of structural materials, which are commonly used in technical practice. 2. methods 2.1. material a structural material pbt filled with 35% of glass fibres (pbt + 35%gf) was chosen as a test subject for this experiment. it is used in numerous technical applications within the automotive industry. the material is labelled pbt v-ptscreatec-b3hzc * m800/25 and was obtained from company pts. it was shipped in the shape of granules that were later enriched by 4 % of cross-linking agent labeled taic (triallyl isocyanurate). exposure of the cross-linking agent to radiation causes a creation of 3d network within the polyamide structure, which is faster than the degradation of the polymer caused by the ionizing radiation. 112 http://dx.doi.org/10.14311/app.2020.27.0112 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 cross-linked polybutylene properties parameters unit pbt + 35%gf injection pressure mpa 80 cooling time s 35 mould temperature °c 85 zone 1 °c 210 zone 2 °c 220 zone 3 °c 235 zone 4 °c 250 table 1. injection molding parameters. 2.2. sample preparation test samples were manufactured by injection molding technology on injection machine arburg allrounder 470h (loßburg, germany). the sample were manufactured according to the standard čsn en iso 179, the dimensions were (80 x 10 x 4) mm. the injection molding process parameters were set in agreement with the material sheet of the tested polyamide. table 1 displays these parameters. the crosslinking causes the connection of polymeric chains to each other, most often using covalent bonds to form the spatial network. test bodies were irradiated under industrial conditions on a commercially available irradiation device in a broader range of radiation doses (0, 66, 132 and 198 kgy) compared to the doses corresponding to the experience in the practice. the irradiation process of the specimens was performed under general conditions (air atmosphere, ambient temperature of 23 °c) just as it is done in engineering practice. one pass in the accelerator exposed the material to 33 kgy of radiation. all samples were irradiated with electron (beta) rays (accelerated electrons a rhodotron r e-beam accelerator, electron energy 10 mev) in the firm bgs beta gamma service gmbh & co, saal am danau germany. 2.3. nano-indentation test nano indentation test was done on most modern indentation device; nano indentation tester (nht3) made by company anton paar (graz, austria). measurements were done by method of instrumented test (dsi depth sensing indentation) according to standard čsn en iso 14577. method used for calculation of mechanical characteristics was oliver pharr. berkovich pyramid was used as measuring tip. the poisson ratio was set to 0.47. it was subtracted from the material sheet for the test material pbt including the 35% glass fibers. device parameters are shown in table 2. the indentation hardness (hit) was calculated as maximum load (fmax) to the projected area of the hardness impression (ap) and the indentation modulus (eit) is calculated from the plane strain modulus (e*) using an estimated sample poisson’s ratio according parameters unit value maximum load mn 50 load/unload speed mn/min 100 holding time s 90 table 2. instrumented test parameters. to [12–15]: hit = fmax ap (1) eit = e ∗ (1 − ν2s ) (2) the indentation creep (cit) was calculated: cit = h2 − h1 h1 · 100, (3) where h1 is the indentation depth at time t1 of reaching the test force (which is kept constant), h2 is the indentation depth at time t2 of holding the constant test force [12–15]. measurement of all above mentioned properties was performed 10 times to ensure statistical correctness. 3. results and discussion measurement done by the instrumented hardness test is based on the principle of immediate detection of indentation in dependence on the applied load (figure 1). the dependence provides the basic information about the behaviour of the tested material, which is subsequently used to calculate the mechanical properties, e. g. indentation hardness and modulus. figure 1. indentation characteristic of tested pbt 35%gf. figure 2 describes the dependence of the indentation depth on time that is used for the calculation of the indentation creep. table 3 shows the average values of the measured parameters. indentation hardness is a basic parameter, which is used to describe the local mechanical properties of the filled pbt surface layer. it is calculated by the oliver and pharr method. the unaltered material displayed 113 m. ovsík, m. staněk, a. dočkal, p. fluxa acta polytechnica parameters unit 0 kgy 66kgy 132kgy 198kgy indentation hardness mpa 65,04 79,35 110,57 116,83 indentation modulus gpa 1,26 1,35 1,83 2,04 indentation creep % 9,80 11,58 9,20 9,83 table 3. indentation parameters. figure 2. indentation characteristic (creep) of tested pbt 35%gf. figure 3. indentation hardness of tested pbt 35%gf. 65 mpa indentation hardness. due to the exposure to the electron radiation, a 3d network was created within the tested pbt, thus significantly increasing the indentation hardness. the indentation hardness measured in test samples irradiated by 66 kgy was 79 mpa, while the test samples irradiated by 198 kgy showed that the indentation hardness rose to 117 mpa. the difference in hardness between the unaltered material and the test sample irradiated by the highest amounts of radiation was 80%. from these measurements, it can be said that due to the filled pbt irradiation, its properties were nearing the properties of the more expensive polymers. furthermore, similar tendencies were measured in the case of indentation modulus. it was 1.3 gpa for the unaltered material and with the exposure of increasing dosages of radiation, the indentation modulus rose up to 2.1 gpa in the material irradiated by 198 kgy. so the improvement was a solid 62%. figure 4. indentation modulus of tested pbt 35%gf. figure 5. indentation creep of tested pbt 35%gf. indentation creep belongs among important local mechanical properties, which are commonly used to describe the behaviour of a specific material exposed to long term stress. as can be seen in figure 5, the irradiation of the samples did not cause any significant changes in the creep behaviour of the filled pbt. the difference between the virgin material and the irradiated pbt was approximately 6%. 4. conclusions this article describes the local nano-mechanical properties of the filled pbt. due to the irradiation by the electron beam the macromolecules were joined within the structure, thus creating a 3d network. this led to a significant improvement of the material properties. the local mechanical 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doi:10.14311/app.2019.22.0139 acta polytechnica ctu proceedings 22:139–144, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app experimental studies of accelerated chloride transport in concrete vojtěch zacharda∗, jiří němeček czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: vojtech.zacharda@fsv.cvut.cz abstract. this contribution deals with the efficiency of electromigration of chlorides used as a repair method for reinforced concrete structures. experimental studies of accelerated chloride transport tests were performed on samples of concrete without chlorides and with admixed sodium chloride during concreting. two concrete types from portland cement characterized with normal and low compressive strengths were studied. the electromigration was applied to penetrate chlorides into the chloride-free sample and for extraction of chlorides from the sample. the effectiveness of the chloride extraction process for rehabilitation of reinforced concrete in terms of lowering the chloride concentration in different concrete types and surface concentration was observed. electrical extraction was found to be effective for lowering of initial chloride concentration by 15-20% after 24 hours. the decrease in surface concentrations was found in the range of 40-50%. the extraction process was found to be feasible and effective for both concrete types. keywords: concrete, electromigration, chlorides, extraction. 1. introduction reinforced concrete structure is exposed to a variety of environmental processes during its lifetime that can cause degradation. the penetration of chlorides into the concrete from deicing salts or in sea splash areas are the most common deterioration processes. chloride ions diffuse through concrete towards steel reinforcement and cause its corrosion. after reaching a critical value of chloride concentration on the rebar surface the corrosion of the steel starts [1–3]. the corrosion is accompanied by an expansion of the oxidic products causing high pressures, concrete cracking and finally spalling of a cover layer. therefore, the resistance of concrete to penetration of chlorides is one of the very important properties for steel-reinforced concretes. since the natural diffusion of ions [4, 5] is a slow process taking years further complicated by concrete aging and chloride binding [6] testing of concrete chloride permeability is often done in an accelerated manner using an electric field [7–9]. both types of tests need to be properly interpreted and diffusion characteristics of concrete can be derived from the tests [9, 10]. electrochemical extraction [7, 8] is one of the solutions that can be used to repair reinforced concrete structure already attacked by chlorides. besides chlorides, other particles with the charge, such as corrosion inhibitors or nanoparticles, can also be transported via electric field [8, 11–13]. the convective flux of ions caused by the electromotive force is much larger that the diffusive flux and leads to radical shortening of the time needed for extraction of chlorides from the structure. in the reality, the electric field is realized between the steel reinforcement and the outer surface electrode (usually stainless steel mesh) placed on the concrete surface. in the lab scale, the tests are performed on concrete samples placed in between the mesh electrodes submerged in a suitable electrolyte. a standard test involves rapid chloride penetration [14] which categorizes samples into levels of vulnerability to the chloride attack. more precise measurements with chloride profile assessment will be presented in this paper. 2. materials two types of concrete mixtures were prepared for this study with low compressive strength (labeled as l) and with normal strength (labeled as n). each type was prepared without chlorides and with 1% of cement weight nacl addition to the mixture prepared from ordinary portland cement cem i-42,5r, sand and natural crushed aggregate. the mixtures composition is shown in the table 1. the concrete was mixed in a 50 l laboratory mixer for 10 minutes. the samples were cylindrical with height of 200 mm and diameter of 100 mm. after casting they were vibrated for approximately 30 seconds. as prevention before water evaporation the specimens were covered with a foil. after 2 days after casting they were unmoulded and stored in water. the resulting 28-days compressive strengths and density of the l samples were 24.3 mpa and 2216 kg/m3 and 50.5 mpa and 2340 kg/m3 for n samples, respectively. for chloride penetration tests the specimens were cut to slices with the thickness of 50 mm. in this study the lower strength concrete represents a weak or damaged material that is worth to be repaired. the different content of cement also play a role 139 http://dx.doi.org/10.14311/app.2019.22.0139 http://ojs.cvut.cz/ojs/index.php/app vojtěch zacharda, jiří němeček acta polytechnica ctu proceedings type of concrete cem i 42,5r sand 0/4 aggregate 4/8 aggregate 8/16 water w/c nacl [kg] [kg] [kg] [kg] [kg] [kg] n 436.4 872.7 290.9 581.8 186 0.43 0 n-1% 436.4 872.7 290.9 581.8 186 0.43 4.36 l 261.8 1150.0 290.9 581.8 210.2 0.8 0 l-1% 261.8 1150.0 290.9 581.8 210.2 0.8 2.62 table 1. concrete mixtures composition, mass per 1 m3. figure 1. an electromigration setup used for accelerated chloride penetration/extraction tests. in chloride binding that is higher in higher strength concrete. 3. methods in order to test the effectiveness of the electromigration process two types of accelerated tests were performed in an migration chamber. first, the chloride penetration test was made on the concrete samples without chlorides (n and l series) and second, the accelerated chloride extraction test was performed on the samples with 1% nacl per cement addition (n-1% and l-1% series). the migration chamber composed of two containers with electrolyte solutions. in case of chloride penetration tests, a 3% nacl solution was used in the compartment with positive electrode and 0.3% naoh solution in the compartment with negative electrode. in case of chloride extraction, both compartments were filed with 0.3% naoh solution and chlorides driven into one of them. a dc power source with constant voltage of 20 v was connected to stainless steel mesh electrodes submerged in the electrolytes. the analyzed samples were inserted between the electrodes and the sample sealed (fig. 1). the accelerated chloride penetration/extraction run for 24 hours. the scheme of tests is shown in fig. 2. before the tests of samples with 1% chloride addition, the original total chloride concentration was analyzed in 20 mm depth steps by drilling of the powder. subsequently, the drilled holes were sealed in the samples and samples analyzed further. after the tests, the concentration was analyzed from powder collected in 5 mm depth steps. always, the powder was mixed with extraction liquid based on potassium dichromate and acetic acid. after at least 24 hours, the solution was analyzed with cl ion selective electrode and the chloride profile constructed. 4. results and discussion 4.1. penetration of chlorides the results of individual accelerated chloride penetration tests were used for calculation of average chloride concentration profile from 3-4 samples as shown in fig. 3 for l samples and in fig. 4 for n samples. the profile exhibits a gradual reduction of chlorides from surface of samples, which was exposed to the 3% nacl solution. the shape of the profile is different in comparison with traditional diffusion tests. here, the chlorides are driven by an electromotive force which is dominant and the profile is a “convective” one with the “middle bulge” caused by the convective force. it can be seen in the figures that the surface concentration after 24 hours exposure is higher by 50% for n samples while the extent of the penetration, i.e. the penetration depth, is larger for l samples. substantial chloride concentration appears within 40 mm of the l sample while only to 10 mm in the case of n sample. the difference in the surface concentration between l and n samples is related to the amount of cement in the mixture and the ability of the concrete to bind chlorides. the binding is higher for n samples which has twice as cement compared to l samples. the density of the n samples is however higher compared to l samples. thus the permeability of the n samples is lower and the penetration depth is substantially lower compared to l samples. 4.2. extraction of chlorides original and after-the-extraction chloride concentration profiles are shown in figs. 5 and 6 for l and n samples, respectively. chlorides are extracted from the sample to the right compartment. the extraction causes decrease in concentration for both sample types. most significantly, the concentration drops down in the surface region (left part of the profile in figs. 5 and 6). the average decrease of chloride concentration was about 15% on l samples and 20% on n samples. the decrease in the surface region was more the 50% for l samples and about 40% for n samples. 140 vol. 22/2019 experimental studies of accelerated chloride transport in concrete (a). (b). figure 2. scheme of accelerated chloride a) penetration test and b) extraction test. figure 3. concentration profile in l samples. the profile is in agreement with the convective transport [10] which leads the chloride ions from left of the samples to the right compartment of the migration chamber. note again, that the transport lasted for 24 hours with voltage of 20 v. further acceleration can be reached by using higher voltages or using larger times. increasing voltage is, however, impractical due to safety reasons and due to joule effect causing heating of the specimen. increasing of the exposure time is, on the other hand, possible and higher efficiency leading to further lowering of the concentration would take place. 5. conclusions the presented paper provides results of accelerated chloride penetration and extraction tests performed on two typical concrete mixtures with different compositions with normal and low compressive strengths. the principle of electromigration was applied to penetrate chlorides into the chloride-free sample and to extract chlorides from the sample with some initial cl concentration. the chloride profiles with concentrations that can be developed after long diffusion tests were gained here in 24 hours while applying safe voltage of 20 v. the chloride concentrations were assessed from concrete core drills using an extraction liquid and ion selective electrode. typical convective profiles with steep concentration change and different surface concentrations were obtained. the different permeability of the samples caused differences in both surface concentrations as well as in penetration depths. the surface concentration was twice as high in n samples compared to l samples. the penetration depth was three to four times larges in l samples compared to n samples. the first phenomenon can be attributed to a larger binding capacity of n samples given by 141 vojtěch zacharda, jiří němeček acta polytechnica ctu proceedings figure 4. concentration profile in n samples. figure 5. concentration profile of l-1% samples before and after extraction. 142 vol. 22/2019 experimental studies of accelerated chloride transport in concrete figure 6. concentration profile of n-1% samples before and after extraction. the twice amount of cement in comparison to l samples. larger penetration depth in l samples is related to their increased permeability given by their lower density and higher porosity. electrical extraction was found to be an effective tool for lowering of initial chloride concentration. 1520% average concentration decrease was reached after 24 hours. the decrease in surface concentrations was even higher (40-50%) which corresponds to the convection of ions out of the samples. the extraction process was found to be feasible and effective for both concrete types. acknowledgements financial support of the czech science foundation (project 16-11879s) and the grant agency of the czech technical university in prague (sgs18/114/ohk1/2t/11) is gratefully acknowledged. references [1] u. angst, ø. vennesland. critical chloride content in reinforced concrete – state of the art. in alexander, et al. (eds.), concrete repair, rehabilitation and retrofitting ii. taylor and francis group, london, 2009. [2] j. cabrera. deterioration of concrete due to reinforcement steel corrosion. cement and concrete composites 18(1):47–59, 1996. doi:10.1016/0958-9465(95)00043-7. [3] k. tuutti. corrosion of steel in concrete. ph.d. thesis, swedish cement and concrete research institute, stockholm, 1982. [4] c. page, n. short, a. el tarras. diffusion of chloride ions in hardened cement pastes. cement and concrete research 11(3):395 – 406, 1981. doi:10.1016/0008-8846(81)90111-3. [5] y. xi, z. p. bažant. modeling chloride penetration in saturated concrete. journal of materials in civil engineering 11(1):58–65, 1999. doi:10.1061/(asce)0899-1561(1999)11:1(58). [6] q. yuan, c. shi, g. de schutter, et al. chloride binding of cement-based materials subjected to external chloride environment – a review. construction and building materials 23:1–13, 2009. doi:10.1016/j.conbuildmat.2008.02.004. [7] m. sánchez, c. alonso. electrochemical chloride removal in reinforced concrete structures: improvement of effectiveness by simultaneous migration of calcium nitrite. construction and building materials 25:873– 878, 2011. doi:10.1016/j.conbuildmat.2010.06.099. [8] h. e. cardenas, l. j. struble. electrokinetic nanoparticle treatment of hardened cement paste for reduction of permeability. journal of materials in civil engineering 18(4), 2006. doi:10.1061/(asce)0899-1561(2006)18:4(554). [9] p. hlaváček, j. němeček. accelerated chloride migration tests in concrete. in proceedings of 22nd international conference engineering mechanics, pp. 213–216. 2016. [10] j. němeček, j. kruis, t. koudelka, t. krejčí. simulation of chloride migration in reinforced concrete. applied mathematics and computation 319:575–585, 2018. doi:10.1016/j.amc.2017.07.029. 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[14] astm c1202-19 standard test method for electrical indication of concrete’s ability to resist chloride ion penetration. standard, american society for testing and materials, west conshohocken. doi:10.1520/c1202-19. 144 http://dx.doi.org/10.4028/www.scientific.net/kem.760.3 http://dx.doi.org/10.1520/c1202-19 acta polytechnica ctu proceedings 22:139–144, 2019 1 introduction 2 materials 3 methods 4 results and discussion 4.1 penetration of chlorides 4.2 extraction of chlorides 5 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.25.0079 acta polytechnica ctu proceedings 25:79–82, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app deformation response of polydimethylsiloxane substrates subjected to uniaxial quasi-static loading vladimír vinarskýa, ∗, fabiana martinoa, giancarlo fortea, jan šleichrtb, václav radab, daniel kytýřb a st. anne’s university hospital brno, international clinical research center, pekařká 53, 656 91 brno, czech republic b czech academy of sciences, institute of theoretical and applied mechanics, prosecká 809/76, 190 00 prague 9, czech republic ∗ corresponding author: vladimir.vinarsky@fnusa.cz abstract. to investigate cellular response of cardiomyocytes to substrate mechanics, biocompatible material with stiffness in physiological range is needed. pdms based material is used for construction of microfluidic organ on chip devices for cell culture due to ease of device preparation, bonding, and possibility of surface functionalization. however it has stiffness orders of magnitude out of physiological range. therefore, we adapted recently available protocol aiming to prepare substrates which offer stiffness in physiological range 5 − 100 kpa using various mixtures of sylgard. an in-house developed loading device with single micron position tracking accuracy and sub-micron position sensitivity was adapted for this experimental campaign. all batches of the samples were subjected to uniaxial loading. during quasi-static experiment the samples were compressed to minimally 40 % deformation. the results are represented in the form of stress-strain curves calculated from the acquired force and displacement data and elastic moduli are estimated. keywords: quasi-static loading, hyperelasticity, polydimethylsiloxane substrates, in-situ loading device. 1. introduction mechanical stimulus emerged as important factor for cell and organ development, homeostasis, and disease in vivo. to investigate effects of static and dynamic mechanical stimulation of in vitro grown cells two different approaches are being currently used. to determine the effect of material stiffness in static conditions polyacrylamide (pa) gels which can be prepared in the range of stiffness reflecting in vivo situation are used. on the other hand, effects of dynamic mechanical stimulation (stretching) are investigated using polydimethylsiloxane (pdms) based microfluidic devices, where cultivation membrane is orders of magnitude stiffer than what can be found in vivo. so far it has been shown that both static and dynamic mechanical stimuli influence cardiomyocyte specification and maturation in vitro. in order to integrate physiological stiffness and possibility to actuate (stretch) the cultivation surface in one device we aimed to prepare pdms based material with stiffness in physiological range, which can be used in pneumatically actuated microfluidic device and test its biocompatibility. as the actuation of the cultivation membrane in general changes material properties we decided to measure the stiffness with tensile strain up to 20 % of original length to mimic its application in microfluidic device. biomechanical analysis of pdms is challenging experimental procedure because of very low material stiffness. there are two basic approaches for this purpose. for the micromechanical tested an atomic force microscopy or dynamic nanoindentation [1] could be used. this experimental procedure is very precise but for bulk material testing macromechanical approach is more appropriate to obtain overall material properties [2]. moreover customised precise macroscopic loading experiments coupled with radiography imaging techniques could provide the information about internal microstructure changes during the loading [3]. for this purpose the material properties of pdms mixtures was obtained employing fully customised experimental system. 2. materials and methods 2.1. sample preparation based on previously published data (see [4]) different ratios of sylgard 184 and sylgard 527 were mixed to obtain stiffness ranging from 5 kpa up to 170 kpa. in order to further decrease stiffness we manipulated the ratio of components a and b of sylgard 527 from 1.2 to 1. the resulting mixtures were degassed, poured in desired vessels and polymerized at 65 ◦c overnight. the bulk material was polymerized in plastic test tubes with nominal inner diameter 12.00 mm. properly polymerized cylinders with diameter 12.00 ± 0.05 mm were carefully removed and carve up to the 79 http://dx.doi.org/10.14311/app.2019.25.0079 http://ojs.cvut.cz/ojs/index.php/app v. vinarský, f. martino, g. forte et al. acta polytechnica ctu proceedings figure 1. visualization of the disassembled loading device with the main components: 1 base plate (for rotary table), 2 loading device frame, 3 motor assembly with harmony gearbox, 4 linear guide assembly, 5 main body cover with clamps, 6 loadcell and loading plate, 7 loading chamber with loading plate samples with height 14 − 16 mm using the microtome blade. plan-parallel faces of the samples are required for relevant testing, therefore only the sampled with height precision ±0.10 mm were selected for the testing. precise height of the samples was derived from loading plates positions at the contact force 0.005 n and 0.010 n for load cells with nominal capacity 1 n and 10 n respectively. for the strain calculation the height was rounded to 50 µm. 2.2. instrumentation estimated high compliance of the material require to use highly precise measuring procedure. therefore, an in-house developed loading device originally intended for 4d microct in detail description in [5] was employed. the device allows to perform analysis of biological samples and artificial product of tissue engineering in controlled ambient conditions. the design of compact table loading device allow to perform the mechanical testing up to 3 kn with single micron position tracking. using the magnetic linear encoder with nominal resolution 250 nm the device operates with sub-micron position sensitivity. using the load cells with appropriate capacity, in this case miniature s beam load cells (lsb200 series, futek inc., usa) with nominal capacity 10 n and 1 n respectively, allows to provide reliable deformation analysis of the materials with elastic modulus in range of single kpa. main components of the loading device are depicted in fig. 1. for this purpose the loading frame was made of a high strength aluminium alloy (en-aw-6082-t6) tube with 60 mm diameter with holes for sample manipulation and optional optical imaging. the loading plates made from the same material with diameter 16 mm ensure the proper support of the faces of the sample under high (≈ 50 %) deformation. the motion of the system and all other peripherals are controlled using a custom developed control software based on the open-source project linuxcnc running on the real-time kernel [6]. the control hardware consists of high-performance motion control components (mesa electronics, u.s.a.), sensor readout electronics (labjack, u.s.a.) and custom electronics. 2.3. experimental procedure to obtain general information about the deformation behaviour of the mixtures and confrontation with findings presented in [4] set of pilot test was performed using 1 to 2 and 1 to 9 (sylgard 184 to sylgard 527) mixtures. it was found that i) the behaviour of the material is fully hyperelastic to minimally 60 % strain ii) the sample is collapsing by sudden outsplashing at level of 70 − 80 % compressive strain iii) the stress-strain behaviour during both loading and unloading phase was strain rate independent on level 0.005 − 0.05 s−1. moreover, the samples due its very low stiffness exhibits high adhesion on the surfaces. it allows to measure tensile behaviour beyond the zero point at unloading of the compressed samples nearly to 10 % tensile strain without any special sample fixation. finally one sample of each type was stored in the fridge with constant temperature of 5 ◦c the other were stored in laboratory temperature of 20 − 25 ◦c. mechanical tests were performed on all samples at as delivered fresh state, after one week and after six weeks. no significant/measurable changes caused aging and/or storage condition were observed. based on the results of pilot test following experimental procedure was applied on all types of the samples. from each type of material three samples were measured. each sample was put to loading chamber and its height was measured by the indicated contact. from this initial point the sample was compressed to the level of 40 % strain on loading rate 20 µm · s−1 (strain rate ≈ 0.015 s−1) and hold on for 5 s at this position. then with the same loading rate the sample was continually unloaded and stretched to 10 %. from this point an other two loading cycled was performed. for the stress-strain analysis and elastic modulus estimation the average unloading curve was calculated. 2.4. data evaluation fully automatic evaluation procedure was developed using matlab (mathworks inc., usa). primary 80 vol. 25/2019 deformation response of pdms substrates figure 2. detail of the compressed pure sylgard 184 sample under 60 % deformation in loading chamber data acquired by control software with sampling frequency 200 hz was imported and force log from load cell and position from encoder was post processed. from the record the unloading phases were extracted and averaged to one unloading force–displacement behaviour. for deformation behavior the stress σ in all experimental analysis was considered as engineering stress obtained using σeng = f ac , (1) where ac is cross-sectional area of the specimen calculated from minimal sample diameter measured before deformation. for the purpose of stress calculations, samples were considered ideally cylindrical, neglecting all geometrical irregularities. engineering strain �eng was calculated by formula �eng = u h0 , (2) where u represents displacement measured by encoder and h0 is height of the sample before the mechanical loading. magnitude of the young’s modulus at level of 0 %, 10 %, and 30 % strain labeled e0.0, e0.1 and e0.3 in stress-strain diagrams (see figs. 3–8) was estimated linear regression from 600 stress–strain values. the linear fit provide less than 0.5 % error. finally the gnuplot (command-line open source program for twoand three-dimensional data and function graphical visualization) code was generated and executed. 3. results material properties and deformation behaviour of twelve different pdms mixtures were derived from set of uniaxial loading tests. each type of material was tested with the batch of the three samples. stress– strain diagram of selected six clinically relevant mixtures together with young’s modulus are depicted and listed in figs. 3–8. light red stripe represents an envelope of measured data from whole batch. dark red curve represents mean values. −2 0 2 4 6 8 10 12 14 16 −0.1 0 0.1 0.2 0.3 0.4 s tr e ss [ kp a ] strain [−] 1:2 (syl184:syl527) e0.0 = 17.0±0.2kpa e0.1 = 27.9±0.2kpa e0.3 = 44.0kpa±0.3 figure 3. deformation behaviour of 1 to 2 (sylgard 184 to sylgard 527) mixture −2 0 2 4 6 8 10 12 −0.1 0 0.1 0.2 0.3 0.4 s tr e ss [ kp a ] strain [−] 1:3 (syl184:syl527) e0.0 = 13.5±0.1kpa e0.1 = 19.8±0.1kpa e0.3 = 31.6±0.2kpa figure 4. deformation behaviour of 1 to 3 (sylgard 184 to sylgard 527) mixture −0.5 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 −0.1 0 0.1 0.2 0.3 0.4 s tr e ss [ kp a ] strain [−] 1:9 (syl184:syl527) e0.0 = 6.4±0.1kpa e0.1 = 8.3±0.2kpa e0.3 = 13.2±0.1kpa figure 5. deformation behaviour of 1 to 9 (sylgard 184 to sylgard 527) mixture −0.5 0 0.5 1 1.5 2 2.5 −0.1 0 0.1 0.2 0.3 0.4 s tr e ss [ kp a ] strain [−] 1:1 (syl527a:syl527b) e0.0 = 3.8±0.1kpa e0.1 = 4.5±0.1kpa e0.3 = 7.5±0.1kpa figure 6. deformation behaviour of 1.0 to 1.0 (sylgard 527a to sylgard 527b) mixture 81 v. vinarský, f. martino, g. forte et al. acta polytechnica ctu proceedings −0.2 0 0.2 0.4 0.6 0.8 1 1.2 −0.1 0 0.1 0.2 0.3 0.4 s tr e ss [ kp a ] strain [−] 1.2:1 (syl527a:syl527b) e0.0 = 1.0±0.1kpa e0.1 = 2.0±0.1kpa e0.3 = 3.6±0.1kpa figure 7. deformation behaviour of 1.2 to 1.0 (sylgard 527a to sylgard 527b) mixture −0.5 0 0.5 1 1.5 2 2.5 −0.1 0 0.1 0.2 0.3 0.4 s tr e ss [ kp a ] strain [−] pure syl527 e0.0 = 3.7±0.1kpa e0.1 = 4.5±0.1kpa e0.3 = 7.0±0.1kpa figure 8. deformation behaviour of pure sylgard 527 4. conclusions to find optimised biocompatible pdms substrates for cardiomyocytes developed for further mechanobiological studies the bulk material with different ratio of sylgard 184, sylgard 527a, and sylgard 527b components. mechanical analysis of soft pdms require employment of custom designed precise loading device measuring the forces and displacement in millinewtons and micrometers respectively. from uniaxial mechanical test deformation behaviour and elastic properties was derived. reliability of the instrumentation and measurement procedure for hyperelastic materials with young’s modulus from single kilopascal corresponding to desired values for further organ-ona-chip systems application [7] was successfully tested . for further analysis focused on dynamic testing simulating physiological processes new voice coil based loading device will be developed for cyclic loading of proposed materials with frequency up to 5 hz. acknowledgements the research has been supported by the european regional development fund in frame of the projects kompetenzzentrum mechanobiologie (atcz133) in the interreg v-a austria czech republic programme. references [1] j. raczkowska, s. prauzner-bechcicki, j. lukes, et al. physico-chemical properties of pdms surfaces suitable as substrates for cell cultures. applied surface science 389:247–254, 2016. doi:10.1016/j.apsusc.2016.07.009. 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[7] s. kratz, c. eilenberger, p. schuller, et al. characterization of four functional biocompatible pressure-sensitive adhesives for rapid prototyping of cell-based lab-on-a-chip and organ-on-a-chip systems. scientific reports 9(1), 2019. doi:10.1038/s41598-019-45633-x. 82 https://doi.org/10.1016/j.apsusc.2016.07.009 https://doi.org/10.14311/app.2019.25.0073 https://doi.org/10.1016/j.matdes.2017.08.036 https://doi.org/10.1371/journal.pone.0051499 https://doi.org/10.1088/1748-0221/13/11/c11021 https://doi.org/10.14311/app.2018.18.0015 https://doi.org/10.1038/s41598-019-45633-x acta polytechnica ctu proceedings 25:79–82, 2019 1 introduction 2 materials and methods 2.1 sample preparation 2.2 instrumentation 2.3 experimental procedure 2.4 data evaluation 3 results 4 conclusions acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0155 acta polytechnica 27(0):155–159, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app nanoindentation of hydrogen enriched zr-1nb zirconium alloy nuclear fuel claddings ondřej liberaa, ∗, patricie halodováa, petra gávelováa, jakub krejčíb a research centre řež, hlavní 130, husinec-řež, czech republic b ujp praha a.s, nad kamínkou 1345, praha-zbraslav, czech republic ∗ corresponding author: ondrej.libera@cvrez.cz abstract. zirconium alloys are being commonly used as a material of choice for nuclear fuel claddings in water cooled nuclear reactors for decades due to their good corrosion resistance and low neutron absorption. however, the increasing operation conditions of next generation nuclear reactors (gen-iv) in terms of higher temperatures, pressures and higher neutron flux requires evaluation of further zr cladding usability. the embrittlement of zr claddings due to hydrogen pickup from reactor coolant is one of the issues for its potential use in gen-iv reactors. nanoindentation is an effective tool for analysis of the change of mechanical properties of hydrogen enriched zr claddings from localised material volume. zirconium alloy zr-1nb (e110) with experimentally induced hydrides was analysed by the means of nanoindentation. zirconium hydrides were formed in the material after exposure in high temperature water autoclave. the optimized methodology of surface preparation suitable for nanoindentation is described and the resulting surface quality is discussed. the nanoindentation measurements were performed as an array of 10x10 indents across areas with hydrides. depth dependent hardness and reduced modulus values measured by nanoindentation were compared between the material with no hydrogen content, low hydrogen content (127 ppm h) and high hydrogen content (397 ppm h). complementary microhardness measurements at hv 0.1 were performed on all materials for bulk material hardness comparison. keywords: nanoindentation, nuclear fuel claddings, zr-1nb. 1. introduction the nuclear fuel cladding serves as the first containment barrier of the fuel pellets in water-cooled nuclear reactors and subsequently in spent fuel storage casks. due to their good mechanical properties, corrosion resistance and low neutron absorption, zirconium alloys are being commercially used as a nuclear fuel cladding material for decades. despite all the benefits, zirconium claddings chemically react with the cooling water in the reactor, which results in creation of hydrides in the material volume [1]. hydrides formed in zirconium claddings affect the mechanical properties of the material, making it prone to cracking due to the internal stresses caused by fuel swelling and the propagation of fission gasses during the nuclear fuel cycle. this process of cracking associated with hydrides in zr alloys is called delayed hydride cracking (dhc) [2]. this phenomenon is one of the key issues with use of zr alloy claddings in gen-iv nuclear reactors. nanoindentation is one of the analytical methods to evaluate the change of mechanical properties before and after hydridation in a very local volume of the material. the mechanical properties of various hydrogen enriched zr claddings [3, 4] and pure zr [5] have already been assessed by nanoindentation measurements in previous studies. however, most of the research in the field was performed on western type on zr claddings such as zircalloy and zr-2.5nb. the amount of available research articles published on nanoindentation of e110 alloy in english is very limited to non-existent, despite the fact that this type of cladding is still being actively used in a large part of the world. in this work, nanoindentation testing was performed on zr-1nb (e110) zirconium alloy nuclear fuel claddings. the tests were conducted on reference and artificially hydrided claddings for future comparison with material irradiated in real operating conditions in pressurised water nuclear reactor (pwr). measurements were performed across the areas containing hydrides. the comparison of mechanical properties between claddings with various amount of hydrogen content is assessed. 2. experimental 2.1. material the material subjected for testing was e110 zirconium alloy cladding tube (zr, 1 wt.% nb, 0.5 wt.% fe, 0.1 wt.% o) in reference state and two hydrogen enriched states. two cladding tubes were exposed to 425°c pressurised deionized water at 10.7 mpa in an autoclave for 63 and 243 days. hydrogen content in samples was evaluated by inert gas fusion (igf) technique and was estimated as 122 ppm and 397 ppm [6]. 155 http://dx.doi.org/10.14311/app.2020.27.0155 http://ojs.cvut.cz/ojs/index.php/app o. libera, p. halodová, p. gávelová, j. krejčí acta polytechnica figure 1. comparison of average hardness values of samples finished by chemo-mechanical polishing for 30 minutes and 1 hour. the final concentration of hydrogen in samples approximately represents the state of cladding during the active use in pwr (122 ppm h) and during dry storage at the end of nuclear fuel cycle (397 ppm h). one milimeter thick rings were cut from each cladding tube on buehler isomet gravitational saw with diamond cutting disc. cut rings were heat embedded in condufast (struers) conductive polymer resin and ground on automatic grinding machine struers tegramin-25 using sandpapers with 500-2000 grits. the samples were then polished on the same machine with 3 and 1 µm diamond suspensions and finished with 0.05 µm colloidal silica suspension ops (struers). the final polishing step was performed on a fine polishing cloth with a mixture of 0.05 µm colloidal silica suspension ops with h2o2 (30%vol.) and hf (40% weight) in 50:10:1 vol. proportion [4]. this chemo-mechanical polishing combination of polishing suspension and etching solution was essential for the preparation of damage-free surface with excellent visibility of the microstructure containing hydrides. commonly used chemical or electrolytic polishing methods for microstructure enhancement and reduction of mechanically hardened layer [7] proved inappropriate for nanoindentation measurements. the inhomogeneous etching rate of zr matrix and hydride phases during these processes introduces increased surface roughness and etch pits, which might compromise the validity of nanoindentation data. 2.2. indentation testing nanoindentation tests were performed on bruker/hysitron ti 950 triboindenter. diamond berkovich type indenter tips were used for all measurements. calibration of tip area function [8] for hardness and reduced modulus calculations were performed on fused silica calibration standard from hysitron (er = 69.6 gpa; h = 9.25 gpa). nanoindentation measurements on e110 samples were performed across the area with visible hydrides as a 10x10 indentation matrix with 20 µm separation between indents. partial unload load function [9] with gradual increase in maximum load up to 10 mn was selected to obtain various depth dependent results for each indentation measurement. hardness and reduced modulus values were calculated from the nanoindentation load displacement curves utilizing the oliver&pharr analytical model [8]. microhardness measurements on struers durascan were performed with diamond tip for hardness comparison at high indentation depths. an array of 10 indents with 100 µm spacing was performed around the hoop direction in the central part of the ring of each sample at hv 0.1 with vickers diamond tip. 3. results and discussion 3.1. sample preparation the duration of chemo-mechanical polishing surface finishing step was found to be crucial for correct nanoindentation measurement on e110 samples. despite the good microstructure visibility after 30 minutes of surface finish, the chemical-mechanical polishing did not completely remove the mechanically hardened layer formed by mechanical grinding and polishing. figure 1 shows the comparison of the measured nanoindentation hardness between chemomechanically polished samples prepared for 30 minutes and 1 hour. the effect of nano-crystallic hardened layer on mechanical properties is best visible on approx. 0.5 gpa increase in overall hardness between hydrogen enriched samples in both states. indentation results presented and discussed later in this work (see section 3.2 and further) are from samples with 1h of surface finish, where the mechanically hardened layer has been reduced by proper sample preparation. 3.2. indentation testing results the area for nanoindentation measurement on 122 ppm h sample was selected to be around the area 156 vol. 27/2020 nanoindentation of hydrogen enriched zr-1nb figure 2. a) light optical micrograph of hydride distribution in a sample with 122 ppm h. b) bse image of indentation matrix in area with hydride on 122 ppm h sample. c) light optical micrograph of hydride distribution in a sample with 397 ppm h. d) bse image of indentation matrix on 397 ppm h sample. figure 3. comparison of average hardness values of reference, 122 ppm and 397 ppm hydrogen enriched e110 samples. e110 reference e110 122 ppm h e110 397 ppm h h 400nm (gpa) 1.56 ± 0.18 1.56 ± 0.12 1.77 ± 0.16 er 400nm (gpa) 72.75 ± 7.63 72.92 ± 2.56 78.14 ± 3.24 hv 0.1 (gpa) 1.44 ± 0.06 1.44 ± 0.05 2.09 ± 0.18 table 1. nanoindentation hardness and reduced modulus at 400 nm indentation depth and microhardness (hv 0.1) results for all samples. with hydride from lom observation shown on figure 2a. at this level of hydrogen pick up the hydrides were formed as individual chains in the hoop direction of the ring. the spacing between individual hydride chains was large, approximately ¼ of the ring width. for this reason, there was only one major hydride chain reached in the nanoindentation matrix (figure 2b). the light and dark stains in bse images represent individual grains with various grain orientations. the hydrides on 397 ppm h sample were much shorter and distributed evenly across the whole area of the ring. the orientation of larger 397 ppm h e110 hydrides was mainly in the hoop direction with minor hydrides in radial direction shown on figure 2c. the 157 o. libera, p. halodová, p. gávelová, j. krejčí acta polytechnica figure 4. comparison of average reduced modulus values of reference, 122 ppm and 397 ppm hydrogen enriched e110 samples. indentation area on 397 ppm h sample was placed in the middle of the ring, covering the representative area with hydrides (figure 2d). the indentation matrix on reference sample was placed in the central part of the ring. the hardness vs. displacement results from nanoindentation measurements plotted in figure 3 show the overall hardness profile across all measured depths for all studied samples. results of the same measurements are shown in figure 4 for reduced modulus values. it is apparent that hardness measurements on all samples were affected by the indentation size effect (ise), causing the increase in hardness values in shallow indentation depths. this effect is well known and unavoidable at these indentation depths [10]. a model for estimation of correct hardness values at low indentation depths was proposed by nix & gao [11], however its application is not in the scope of this work. the increased reduced modulus at low indentation depths (<200 nm) might be caused by a thin oxide layer formed after sample preparation. this assumption has to be confirmed by further investigation of the surface after the preparation process. an increase of hardness and reduced modulus is visible at 397 ppm h sample. a dense mesh of hydrides in the material directly affects the mechanical properties of the whole material. measured mechanical properties of 122 ppm h sample are nearly identical to values from the reference sample. it can be assumed that the scarce presence of hydrides in 122 ppm h sample affects the mechanical properties of the whole material just slightly. however, a closer examination of the mechanical properties near the hydride has to be performed to confirm this assumption. comparison of hardness measured by nanoindentation and hv 0.1 microhardness is shown in table 1. the value of 400 nm was selected as a stable nanoindentation depth without the influence of the ise for comparison. a good agreement of measured values is shown for reference e110 and 122 ppm h sample. the microhardness results for 397 ppm h show approx. 0.3 gpa higher value than the nanoindentation measurement. the reason behind this difference is a matter of further research. 4. conclusions optimized sample preparation method was used for fine surface preparation of e110 samples for nanoindentation measurements. samples of zr-1nb e110 nuclear cladding tubes in reference, 122 ppm and 397 ppm hydrogen enriched states were measured by nanoindentation. it was shown that hardness and reduced modulus values of the 122 ppm h sample were very close to the reference sample. this result was confirmed by comparison to microhardness testing results. a further research needs to be performed closer to hydride chains on 122 ppm h sample to confirm the preservation of mechanical properties despite the hydrogen build-up. the measurements on 397 ppm h sample show an increase of hardness by approx. 13 % and reduced modulus increase by 8 %. the measured microhardness value was 18 % higher than nanoindentation hardness measured at 400 nm. this phenomenon is a subject of further research. acknowledgements the presented work was financially supported by the ministry of education, youth and sport czech republic project lq1603 (research for susen). this work has been realized within the susen project (established in the framework of the european regional development fund (erdf) in project cz.1.05/2.1.00/03.0108). references [1] c. coleman, d. hardie. the hydrogen embrittlement of α-zirconium-a review. journal of the less common metals 11(3):168 – 185, 1966. doi:10.1016/0022-5088(66)90003-8. [2] delayed hydride cracking in zirconium alloys in pressure tube nuclear reactors. final report of a coordinated research project 1998-2002. tech. rep. 41, international atomic energy agency, nuclear power technology development section, vienna (austria), 2004. 158 https://doi.org/10.1016/0022-5088(66)90003-8 vol. 27/2020 nanoindentation of hydrogen enriched zr-1nb [3] k. kese, p. olsson, a.-m. a. holston, e. broitman. high temperature nanoindentation hardness and young’s modulus measurement in a neutron-irradiated fuel cladding material. journal of nuclear materials 487:113 – 120, 2017. doi:10.1016/j.jnucmat.2017.02.014. [4] a. rico, m. martin-rengel, j. ruiz-hervias, et al. nanoindentation measurements of the mechanical properties of zirconium matrix and hydrides in unirradiated pre-hydrided nuclear fuel cladding. journal of nuclear materials 452(1):69 – 76, 2014. doi:10.1016/j.jnucmat.2014.04.045. [5] m. kuroda, d. setoyama, m. uno, s. yamanaka. nanoindentation studies of zirconium hydride. journal of alloys and compounds 368(1):211 – 214, 2004. doi:10.1016/j.jallcom.2003.08.094. [6] b. d. holt, h. t. goodspeed. determination of nitrogen, oxygen, and hydrogen in metals by inert gas fusion. a manometric method. analytical chemistry 35(10):1510–1513, 1963. doi:10.1021/ac60203a023. [7] g. finch, a. quarrell. the beilby layer. nature publishing group 137(3467):516–519, 1936. doi:10.1038/137516a0. [8] w. oliver, g. pharr. an improved technique for determining hardness and elastic modulus using load and displacement sensing indentation experiments. journal of materials research 7(6):1564–1583, 1992. doi:10.1557/jmr.1992.1564. [9] a. c. fischer-cripps. nanoindentation. springer, new york, ny, 3rd edn., 2011. doi:10.1007/978-1-4419-9872-9. [10] y. milman, a. golubenko, s. dub. indentation size effect in nanohardness. acta materialia 59(20):7480 – 7487, 2011. doi:10.1016/j.actamat.2011.08.027. [11] w. d. nix, h. gao. indentation size effects in crystalline materials: a law for strain gradient plasticity. journal of the mechanics and physics of solids 46(3):411 – 425, 1998. doi:10.1016/s0022-5096(97)00086-0. 159 https://doi.org/10.1016/j.jnucmat.2017.02.014 https://doi.org/10.1016/j.jnucmat.2014.04.045 https://doi.org/10.1016/j.jallcom.2003.08.094 https://doi.org/10.1021/ac60203a023 https://doi.org/10.1038/137516a0 https://doi.org/10.1557/jmr.1992.1564 https://doi.org/10.1007/978-1-4419-9872-9 https://doi.org/10.1016/j.actamat.2011.08.027 https://doi.org/10.1016/s0022-5096(97)00086-0 acta polytechnica 27(0):155–159, 2020 1 introduction 2 experimental 2.1 material 2.2 indentation testing 3 results and discussion 3.1 sample preparation 3.2 indentation testing results 4 conclusions acknowledgements references doi:10.14311/app.2021.30.0076 acta polytechnica ctu proceedings 30:76–80, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app increasing the biological resistance of gypsum by means of silver, calcium and copper ions zdeněk prošeka, b, ∗, hana sekavováa, pavla ryparováa a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: zdenek.prosek@fsv.cvut.cz abstract. the article focuses on the effects of additives increasing biological resistance of gypsum against degradation of the material. in this study we use silver nitrate, calcium oxide and copper sulphate as an additive that improves biological resistance of gypsum composites. the amount of addition was calculated so that the amount of ions was 0.1, 0.3 and 0.5 wt. %. microbial degradation was determined as the degree of resistance of materials to the occurrence of bacteria and molds by optical and electron microscope. the results obtained from these samples were compared with references materials created in the laboratory and available on today’s market. the results show a positive effect of silver and copper ions on the bioresistance of gypsum. keywords: biological resistance, calcium oxide, copper sulphate, gypsum, silver nitrate. 1. introduction gypsum is a widely used material in construction, especially for the production of gypsum plaster, wallboard, gypsum blocks [1]. gypsum is one of the longest used binders, it was already known in egypt or assyria. it has been used in europe since the early middle ages [2]. gypsum composites are more environmentally friendly and production is faster than cement composites. mainly due to lower greenhouse gas emissions and sustainability [3]. gypsum is produced by thermal decomposition of caso4 · 2h2o in a so-called boiling process at 130 − 150 ◦c [4]. the material for the production of gypsum can be natural or also waste after flue gas desulphurisation from thermal power plants (industry gypsum). the use of gypsum composites in construction is significantly limited because they have low resistance to damage of moisture. the moisture content of the material adversely affects the formation of bacteria, molds and the multiplication of other microorganisms [5]. molds are microorganisms that commonly occur on the inner and outer surfaces of building materials and can cause sick building syndrome, which limits or completely restrict the use of the building [6]. for this reason, it is very important to create biological resistance materials that resist the formation of mold on their surface and inside the material. one of the ways to increase the biological resistance of gypsum is to prevent the access of water or moisture. without the presence of water, mold and bacteria colonies cannot form. reduced water absorption is achieved by the use of hydrophobic additives such as waxes [7], saponifiers [8] and emulsifiers [9]. this method of protecting gypsum is very effective, but prevents possible future recycling of gypsum. during recycling, the hydrophobicizing additive cannot be separated from the gypsum and thus the calcined gypsum can no longer be mixed properly with the water and thus prevents it from solidifying properly [10]. another option is to create an environment that does not allow the growth of molds and bacteria. for example, it is a very alkaline environment that restricts the growth of bacteria and molds [11]. the last possibility to improve the biological resistance of gypsum composites is by adding antibacterial components (such as silver [12], titanium dioxide [13], copper [14]). this article deals with the direct use of silver and copper ions as additives that increase biological resistance. when penicillin was discovered, silver and other biocidal materials were almost abandoned. but today, with the emergence of antibiotic-resistant strains, it has gained new interest [15]. silver has been reported in several studies [16, 17] as an effective bactericidal antibacterial agent. gypsum composite with antibacterial components is also being developed and investigated [5], but some of its properties are still unknown. in an experimental study, the effect of silver ions on the mechanical properties of the gypsum composite was investigated and it was found that: the introduction of silver ions results in a less compact structure of hardened gypsum samples, leading to a 20% deterioration in flexural strength [18]. work dealing with other additives is completely missing. 2. materials and samples in our study of gypsum biodegradation, three reference materials were used: ref 1 was labora76 http://dx.doi.org/10.14311/app.2021.30.0076 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 increasing the biological resistance of gypsum set gypsum [g] admixtures water [g] ions and their content ref 1 500 320 ag 0.1 500 agno3 – 1.33 g 320 ag – 0.1 wt. % ag 0.3 500 agno3 – 4.02 g 320 ag – 0.3 wt. % ag 0.5 500 agno3 – 6.69 g 320 ag – 0.5 wt. % ca 0.1 500 cao – 1.19 g 320 ca – 0.1 wt. % ca 0.3 500 cao – 3.57 g 320 ca – 0.3 wt. % ca 0.5 500 cao – 5.94 g 320 ca – 0.5 wt. % cu 0.1 500 cuso4 – 2.14 g 320 cu – 0.1 wt. % cu 0.3 500 cuso4 – 6.41 g 320 cu – 0.3 wt. % cu 0.5 500 cuso4 – 10.68 g 320 cu – 0.5 wt. % table 1. composition of tested sets. 100 µm figure 1. image from electron microscope, bse detector, magnified 400×. tory produced according to table 1, ref 2 is made of gypsum slurry for plasterboards (marked h2) designation with reduced water absorption and thus higher bioresistance, according to the standard čsn en 520 + a1 [19] and ref 3 is made of gypsum slurry for plasterboards marked a (common use), according to standard čsn en 520 + a1 [19]. ref. 2 and 3 were used to compare with current materials on today’s market. the resistance to biodegradation of the laboratory mixture was increased by ions of silver, calcium and copper. silver ions were obtained from the addition of silver nitrate (agno3), calcium ions from calcium oxide (cao) and copper ions from copper sulphate (cuso4). the amount of addition was calculated so that the amount of ions was 0.1, 0.3 and 0.5 wt. % (table 1). silver and copper ions were chosen because of their biocidal properties and calcium due to increased alkalinity of the environment which prevents the growth of molds and bacteria. five samples were made from each set according to table 1. the samples were stored for 9 days in a laboratory environment at temperature 22 ± 3 ◦c and relative humidity 50 ± 5 %. subsequently, the samples were tested. figure 2. sample placed on agar in petri dish after 41 days. in all cases, gypsum from the knauf praha company was used, originating from the industry gypsum of the počerady power plant. figure 1 show the typical character of gypsum grains, which are composed of only one phase and have an elongated, onesided oriented shape. the results of the eds analysis show that the gypsum grain is composed of calcium 19 wt. %, sulfur 17 wt. % and oxygen 64 wt. %. when converted to atomic percent and using stoichiometry, it can be stated that it is pure gypsum. 3. experimental methods microbial degradation was performed according to the standard iso en 846 [20], which determines the degree of resistance of materials to the occurrence of molds or determines their resistance to biodegradation. the produced samples had a size of 1 × 1 × 1 cm. sterilization was performed using uv radiation, which damaged microorganisms from the surface while maintaining the material and chemical parameters of individual samples. other types of possible sterilization, apart from radioactivity, degrade the original properties of the materials and were therefore not used. samples were inoculated with a 77 z. prošek, h. sekavová, p. ryparová acta polytechnica ctu proceedings ref. 1 ref. 2 ref. 3 10 mmmold colonies figure 3. surface of reference samples after 41 days of exposure to molds and bacteria. ag 0.1 ca 0.1 cu 0.1 ag 0.3 ca 0.3 cu 0.3 ag 0.5 ca 0.5 cu 0.5 mold colonies copper inclusion 0.18 mm figure 4. optical microscope images of tested materials. magnified 120×. consortium of molds and then plated on agar (czapek dox) at 25 ± 3 ◦c and monitored for 41 days (fig. 2). the consortium of molds contained microorganisms from the czech collection of microorganisms: paecilomyces variotti, chaetomyum globosum, penicillium funiculosum, aspergillus terreus, penicillium brevicompactum. the resistance of the materials was evaluated visually and monitored by optical microscope. optical microscopy was performed using a zoom microscope with 3d imaging capabilities. this is a device from zeiss axio zoom.v16. clusters of additive ions were discovered during optical microscopy. for this reason, optical microscopy has been supplemented by electron microscopy. a scanning electron microscope with a schottky cathode feg sem merlin zeiss was used for electron microscopy, which is located in the laboratory of electron microscopy and micro-analysis at the university center for energy efficient buildings in buštěhrad. quantitative and qualitative analysis of the chemical composition of the samples was performed by x-ray microanalysis directly with an energy dispersive spectrometer (eds) from oxford instruments. during the anal78 vol. 30/2021 increasing the biological resistance of gypsum 500 µm ag 0.1 ca 0.1 cu 0.1 ag 0.5 ca 0.5 cu 0.5 figure 5. image from electron microscope, bse detector, magnified 200× with eds element maps (yellow – silver, white – calcium and blue – copper). ysis, the microscope settings were as follows: accelerating voltage 20 kv, current 2 na, resolution 1024 px, work distance 8.5 mm and acquisition time 60 s. due to the higher electric current required for elemental analysis, 3 nm of platinum were sputtering on surface of samples. 4. results and discussion figure 3 shows a comparison of reference materials. the ref. 3 sample had almost the entire sides covered, and mold colonies were already present on the surface on day 22. the coverage rate did not change much at the end of the experiment, but the molds began to fructify. samples ref. 1 and 2 had similar results and were generally resistant to the occurrence of molds, after 41 days only isolated colonies of molds were visible on the surface, the sides of the samples were only partially covered, approximately 40 %. the laboratory mixture ref. 1 achieved the same results as gypsum slurry with lower water absorption (ref. 2) due to the fact that it did not contain paper pulp and other materials contained in gypsum slurry for plasterboards which reduces bio-resistance. in the case of added ions, all variants were highly resistant to biodegradation. for this reason, optical microscopy was performed. figure 4 shows the surfaces of the tested materials by optical microscope. the results show a positive effect of copper and silver, where after 41 days no microorganisms formed on the surface of the materials. in the case of calcium ions, a mold colony can be found on the surface of all materials tested. silver and copper ions 100% increase the bioresistance of gypsum even in the lowest tested amount of ions. in addition, it can be seen that with a higher concentration of copper ions, clusters occur. the individual ion clusters can be seen in figure 5 from electron microscopy. the images were taken using a bse detector and show the phase contrast. in addition, the left part of the images is supplemented by an element map, where elements were selected for individual ions. silver is shown in yellow, calcium in white and copper in blue. the results show that higher concentrations (0.5 wt. %) of silver and copper ions form clusters. in the case of a concentration of 0.1 wt. %, the ions are homogeneously distributed throughout the structure. individual cao grains can be seen in the calcium sample (ca 0.1 and ca 0.5). the results of the influence of silver and copper ions are in agreement with other authors who have used applications for other building materials such as lime and lime-cement plasters [21]. the formation of clusters explains the decreasing mechanical properties of gypsum with a high content of silver ions, which are described by other authors [18]. 5. conclusions this work was focused on the influence of different type of admixtures on biological resistance of gypsum. the researched samples were composed of gypsum from knauf praha (počerady) and various types of biocidal material. silver, calcium and copper ions in three different concentrations (0.1, 0.3 and 0.5 wt. %) 79 z. prošek, h. sekavová, p. ryparová acta polytechnica ctu proceedings were used as biocidal material. based on the results, it can be concluded that: • the addition of silver, calcium and copper ions has increased the resistance to gypsum biodegradation and the materials are comparable to those on the market today. • higher ion concentration (0.5 wt. %) form clumps and inhomogeneity of materials. • silver and copper ions were most effective, where no colony of molds and bacteria formed after 41 days for all ion concentrations. future research will focus on finding a minimum concentration of silver ions in gypsum that will be resistant to biodegradation. next research will not deal with calcium ions, due to lower efficiency and copper ions, due to toxicity and current legislation [22]. acknowledgements this paper was financially supported by the ta cr research a project under the number tj02000235 and by czech technical university in prague under no. sgs project sgs19/148/ohk1/3t/11. the authors also thank knauf, praha ltd. for the supplied samples and materials and the center for nanotechnology in civil engineering at the faculty of civil engineering, czech technical university in prague. references [1] m. doleželová, a. vimmrová. moisture influence on compressive strength of ternary gypsum-based binders. in aip conference proceedings, vol. 1863, p. 290003. aip publishing llc, 2017. doi:10.1063/1.4992440. 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[19] čsn en 520+a1. sádrokartonové desky definice, požadavky a zkušební metody. praha: úřad pro technickou normalizaci, metrologii a státní zkušebnictví . [20] čsn en iso 846. plasty hodnocení působení mikroorganismů. praha: úřad pro technickou normalizaci, metrologii a státní zkušebnictví 2019. [21] a. jerónimo, a. camões, b. aguiar, n. lima. hydraulic lime mortars with antifungal properties. applied surface science 483:1192–1198, 2019. doi:10.1016/j.apsusc.2019.03.156. [22] n. a. ashford, c. c. caldart. environmental law, policy, and economics: reclaiming the environmental agenda. mit press, 2008. 80 https://doi.org/10.1063/1.4992440 https://doi.org/10.1007/bf01063410 https://doi.org/10.1016/b-978-0-08-100370-1.00025-1 https://doi.org/10.14311/app.2019.21.0016 https://doi.org/10.1080/00038628.2018.1461060 https://doi.org/10.1177/00220345580370022301 https://doi.org/10.1016/j.conbuildmat.2019.06.227 https://doi.org/10.14311/app.2020.26.0081 https://doi.org/10.1016/j.biotechadv.2008.09.002 https://doi.org/10.2217/fmb.11.78 https://doi.org/10.1016/j.ijantimicag.2008.12.004 https://doi.org/10.1093/jac/dkm006 https://doi.org/10.1007/s13762-018-1773-0 https://doi.org/10.1016/j.apsusc.2019.03.156 https://doi.org/10.14311/app.2022.33.0284 acta polytechnica ctu proceedings 33:284–289, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague lower bound analysis for shear assessment of full-scale rc girders subjected to axial tension takeru kanazawa hokkai gakuen university, 1-1, minami 26, nishi 11, tyuo-ku, sapppro, hokkaido, japan correspondence: t-kanazawa@hgu.jp abstract. axial tension force exerted as a result of a temperature change or shrinkage can cause the collapse of rc structural members. design code provisions and analytical models such as modified compression field theory (mcft) yield reasonable estimates of shear strength of rc beams subjected to axial tension. nevertheless, their semi-empirical nature is not necessarily appropriate for shear assessment of existing rc structural members. the extra conservativeness and empirically determined parameters might require unnecessary maintenance work. a generalised model with rigorous formulation must be developed. this paper presents a purely theoretical model to predict the shear strength of rc beams under axial tension based on limit analysis. without regressive functions and empirical functions, lower bound analysis enables shear strength derivation when the force equilibrium and strain compatibility are satisfied. accuracy of the analysis was verified by comparison of its predictions with three experimental shear strengths of full-scale rc girders. an equal level of accuracy was observed between the analytical solutions and mcft-based predictions. keywords: axial tension, full-scale rc girders, lower bound analysis, shear assessment. 1. introduction the sustainability of rc structures is an important issue worldwide because of their exposure to a variety of loading with environmental and mechanical actions. the partial collapse of the wilkins air force depot in 1955 [1] spurred the refinement of design practices for shear under axial tension caused by a temperature change or shrinkage. the brittle collapse was reproduced in laboratory investigation [1]. results demonstrated that tensile stress of approximately 1.4 n/mm2 can reduce shear strength by 50% [2]. design code provisions have difficulty in predicting such detrimental effects of axial tension because of empirically determined parameters with safety factors. in this respect, several well-established theories for shear [3–5] are based not only on the force equilibrium, but on the failure kinematics. the former and latter respectively correspond to lower and upper bound solutions. the lower bound analysis is the basis of the shear design expression in most international codes. this lower bound is the reason underlying their inherent conservativeness, which is appropriate for designing new structures. for structural assessment of existing rc members, however, such conservativeness is of less importance than determining the actual bearing capacity [6]. among the currently available theories, the modified compression field theory (mcft) has provided a consistent framework to consider the influence of axial force and to facilitate its incorporation into various international codes and specifications. nevertheless, the mcft predictions might elicit some discrepancy with several test results of rc beams under axial tension [2]. current understanding should be supplemented theoretically for rational shear assessment because the limitations of existing test results include laboratory-scale specimens without web reinforcement [7], despite recent progress [8, 9]. this paper presents a rational analysis based on the lower bound theorem. the analytical structure is based on the combined upper and lower bound analysis [10], which predicted the shear strengths of rc beams accurately with no axial force and no web reinforcement. the present work extends their original theory to consider the effectiveness of web reinforcement and the negative effects of axial tension. rigorous formulation for shear failure is presented, including shear resisting mechanisms such as aggregate interlock and dowel action of longitudinal reinforcement under both force equilibrium and strain compatibility. the developed analysis is validated by comparison of its solutions with mcft-based predictions and test results of three full-scale rc girders subjected to axial tension [7]. the notation used for this study is summarised in the appendix. 2. model formulation 2.1. analytical framework in the theoretical framework of limit analysis [11], the exact solution lies between lower and upper bound solutions, for which the former and latter respectively satisfy force equilibrium and compatible patterns of failure, in addition to yield criteria. one therefore finds the exact solution by maximising the lower bound solutions. to derive shear strength with 284 https://doi.org/10.14311/app.2022.33.0284 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 analysis for shear assessment of full-scale rc girders figure 1. constitutive law of concrete. figure 2. constitutive law of steel reinforcement. reasonable simplification, the following assumptions are introduced. • concrete and steel reinforcement are in a plane stress state. • perfect bonding exists between steel reinforcement and concrete. • concrete strength equal to over a compression zone; concrete behaves as a rigid perfectly plastic material, as presented in figure 2. • steel reinforcement behave as an elastic perfectly plastic material, as shown in figure 2. it is noteworthy that the physical quantities such as forces, stresses and strains in compression are negative. those denoted with a prime have positive compression. 2.2. force equilibrium figure 3 shows the free-body diagram with the shear resisting components against the external shear and axial tension force, respectively denoted as v and n. this equilibrium condition yields the following expressions, respectively, for vertical force, horizontal force, and moment. v − vc − fd − fw − b ! t f sin θ dt = 0 (1) figure 3. free-body diagram. figure 4. critical crack shape. ! h αh bf ′c dy − fs1 − f ′ s2 + n + b ! t f cos θ dt = 0 (2) [γa + cot β (h − d1)] fd + (h − d1) fs1 − (h − d2) f ′s2 + a (1 + γ) 2 fw + avc − h 2 n − h (1 + γ) 2 ! h αh bf ′c dy + b "! t f sin θ x dt − ! t f cos θ y dt # = 0 (3) therein, $ t dt denotes the path integral from point a to b in figure 3. one can replace these integrations with the following integrations with respect to x and y. ! t cos θ dt = ! a γa dx, ! t sin θ dt = ! αh 0 dy (4) mathematical treatment of the path integral requires an expression for the critical crack shape shown in figure 4 on the x − y coordinates. y2 = α2h2 (x − aγ) a (1 − γ) (5) dividing equations 1, 2 and 3 respectively by bd1f ′c and abd1f ′c, one obtains non-dimensionalised expressions as presented below. 285 takeru kanazawa acta polytechnica ctu proceedings figure 5. strain compatibility. v = (1 − α) τc + φs1τs + ωφswσw + α µ f (6) 1 − α µ − φs1σs1 + φs2σs2 + n + λ (1 − γ) f = 0 (7) % γ + cot β " ζ − 1 λ #& τsφs1 + " ζ − 1 λ # φs1σs1− " ζ − δ λ # φs2σs2 + 1 + γ 2 ωφwσw + (1 − α) τc − ζ 2 n − ζ 2µ ' 1 − α2 ( − α 3µ (1 − 4γ) f = 0 (8) in those equations, the non-dimensional static variables are defined as shown below. ) **********+ **********, v = v bd1f ′c , σs1 = fs1 as1fs1 , σs2 = f ′s2 as2fs2 , σw = fw aswfsw , τc = vc bd1f ′c (1 − α) , τs = fd as1fs1 , n = n bd1f ′c , f = f f ′c . (9) equation 6 is the objective function in this study. the non-dimensional shear strength of v is maximised with respect to the static variables. 2.3. compatibility conditions though lower bound solutions do not require compatibility conditions, an inequality constraint for aggregate interlock ' f ( can be derived from strain compatibility. figure 5 presents the strain compatibility on the cracked section. the assumption of bernoullieuler beam (plane sections remain planar) yields the linear distribution of longitudinal strain. those of bottom and top reinforcement are obtained respectively as shown below. εs1 = d1 − (1 − α) h (1 − α) h ε′cu = 0.0035 µ 1 − α − 1 . (10) εs2 = (1 − α) h − d2 (1 − α) h ε′cu = 0.0035 δµ 1 − α − 1 . (11) the neutral axis depth under the effect of axial tension is calculable by arranging equation 6 with respect to k = (1 − α)h/d1 as presented in the equation below. k = ε′cues f ′c / (ρs1 + δρs2) µ 1 − α − (ρs1 + ρs2) 0 − f λ (1 − γ) − n (12) equation 12 gives an upper limit of the aggregate interlock in maximisation because k must be positive according to its definition. 2.4. yield criterion and solution technique the following von mises yield criterion is introduced between τs and σs1. 3τ 2s + σ 2 s1 = 1 (13) although hweé et al. [10] used another yield criterion for concrete, it need not be addressed in this study because the rigid perfectly plastic behaviour is assumed. the complete set of equations enables the maximisation of the objective function of equation 6 with respect to the six static variables ' σs1, σs2, σw, τc, τs, and f ( . the possible ranges of those variables are presented as the following. σs1, σs2, σw ∈ 〈0, 1〉 τc, τs, f ∈ 〈0, ∞〉 (14) all of those expressions must be positive. the upper limits of σs1, σs2, and σw and correspond with yielding of reinforcement. although no upper limit is imposed on the others, equation 7, 8, 12, and 13 render them physically admissible values. maximisation is performed using the optimisation algorithm of matlab [12], under the equality and inequality constraints listed in table 1. all possible 286 vol. 33/2022 analysis for shear assessment of full-scale rc girders objective function equation 6 equality constraints (force equilibrium) equation 7 and 8 equality constraints (yield criterion) equation 13 inequality constraints (neutral axis depth) equation 12 ≥ 0 inequality constraints (yield criterion) 3τ 2s + σ2s1 ≤ 1 table 1. list of constraints for optimisation. figure 6. cross-sectional details of inverted-t beams (unit: mm). figure 7. shear strength dependence on applied axial force. values of v are calculated within the following ranges of kinematic variables. 1 − µ ≤ α ≤ 1 αh a (1 − γ) ≤ β ≤ ∞ 0 ≤ γ ≤ 1 (15) the lower limits of and stem respectively from the assumption that the longitudinal strain of bottom reinforcement in equation 10 must be positive, and that the critical crack is directed upwards at y = 0 (figure 3). to obtain more optimal solution, the maximised solutions are then minimised with kinematic variables. the partial derivative of equation (6) with respect to determines whether the objective function is monotonically decreasing or increasing. the minimised shear strengths are calculable as presented below. ∂v ∂α = − τc + f µ (16) ) *******************+ *******************, vana = φs1τs + ωφswσw + f µ " if ∂v ∂α < 0, αmax = 1 gives vana # vana = µτc + φs1τs + ωφswσw + 1 − µ µ f " if ∂v ∂α > 0, αmax = 1 − µ gives vana # (17) 3. results and discussion 3.1. experiment results the analytical solutions have been compared with test results obtained for full-scale rc girders [7]. the studied girders had the inverted-t configuration (figure 6) to support the deck in flexural tension. figure 7 presents the detrimental effect of axial tension on shear strength. the axial force of 900 kn decreases the shear strength by 14%. the axial force magnitudes of 3-200-p and 4-300-p were found respectively to represent shrinkage-induced stress only, and both the shrinkage and temperature-induced stress. it is noteworthy that smith et al. [7] tested seven specimens in all. the three specimens shown in figure 7 were extracted for model validation because the others included the use of cut-off bars and epoxy injected specimen, of which the contributions were not formulated. 3.2. model validation to calculate the shear strength by the developed analysis, static and kinematic constants were determined as listed respectively in tables 2 and 3. these values were calculable using information referred from reports of the related literature [7]. the rectangular area of bd1 in figure 6 was used for calculation because the contributions of the flanges under flexural tension are negligible. in addition, the values of ε′cu 287 takeru kanazawa acta polytechnica ctu proceedings n = n bd1f ′c φs1 = ρs1fs1 f ′c φs2 = ρs2fs2 f ′c φw = ρswfsw f ′c 1-0-p − 0.221 0.110 0.064 3-200-p 0.083 0.269 0.135 0.079 4-300-p 0.128 0.267 0.133 0.081 table 2. non-dimensional static constants. λ = a d1 ζ = h a µ = d1 h σ = d1 d2 ω = s d1 1-0-p 3-200-p 2.60 0.407 0.943 0.066 0.265 4-300-p table 3. non-dimensional kinematic constants. vexp vexp vana vaasht o vana vexp vaasht o vexp + vdl (kn) (−) (−) (kn) (−) (−) 1-0-p 902 0.0773 0.0809 810 1.05 0.88 3-200-p 780 0.0814 0.0879 743 1.08 0.93 4-300-p 783 0.0835 0.0847 712 1.01 0.89 table 4. accuracy of the developed analysis. and es in equation (12) are assumed respectively as 0.0035 and 200 kn/mm2 because the original data were not found in the literature [7]. table 4 presents comparisons among test results (vexp), present analysis (vana), and mcft-based prediction by aashto lrfd bridge design specification [13]. the analytical solutions show adequate accuracy (vana/vexp) with specimens under axial tension (3-200-p and 4-300-p) and without axial tension (1-0-p). in all cases, the static variable of σw was equal to one when the optimisation terminated. this value represents the yielding of web reinforcement, which was observed in the experiments as well. furthermore, results demonstrate better agreement than the mcft-based predictions, except for those for specimen 3-200-p. although the analytical predictions of 3-200-p and 3-400-p show a consistent decrease of shear strength with the increase of applied axial force, no such tendency was observed in the experiment as presented in figure 7. further investigation about this discrepancy will be reported in another paper with a broad range of experimental data. 4. conclusions shear strengths under the effect of axial tension loading were derived analytically based on the lower bound theorem. the theoretical basis treated the shear strength derivation as an optimisation problem under equality and inequality constraints obtained from force equilibrium, strain compatibility and yield criteria. model validation showed good correspondence between experimental and analytical shear strengths without any regression functions and empirically determined parameters. the accuracy was an equal level with mcft-based predictions. the generality of this formulation enables the consistent treatment of axial compression, and gives quantitative results of each shear component such as aggregate interlock of cracked concrete and dowel action of longitudinal reinforcement. although further verification by comparison with a number of test results is necessary, the developed analysis might contribute to the provision of comprehensive understanding of shear behaviour under axial force, together with the mcft [2] and other well-established theories [3–5]. research is currently underway to extend this theory to axially compressed and tensioned beams without web reinforcement. list of symbols f ′c concrete compressive strength v external shear force n external force of axial tension vc shear component of uncracked concrete fw shear component of web reinforcement f aggregate interlocking force per unit area θ inclination angle of critical crack to x-axis fd dowel force of bottom reinforcement fs1, fs2 longitudinal forces of bottom and top reinforcement, respectively α critical crack height divided by total depth 288 vol. 33/2022 analysis for shear assessment of full-scale rc girders β slope of critical crack at point b γ longitudinal distance between support and point b divided by shear span a shear span b web width d1 effective depth h total depth d2 distance between top reinforcement and upper perimeter of beams s spacing between adjacent web reinforcement t coordinate axis along critical crack ω spacing between adjacent web reinforcement divided by effective depth µ effective depth divided by total depth λ slenderness (shear span-to-depth) ratio ζ total depth divided by shear span δ d1 divided by d2 k neutral axis depth divided by effective depth ε′cu ultimate compressive strain of concrete (0.0035) es modulus of elasticity of longitudinal reinforcement (200 gpa) ρs1, ρs2, ρsw reinforcement ratios of bottom, top and web reinforcement, respectively as1, as2, asw areas of bottom, top and web reinforcement, respectively fs1, fs2, fsw yield strengths of bottom, top and web reinforcement, respectively εs1, εs2, εsw longitudinal strains of bottom and top reinforcement, respectively φs1, φs2, φsw reinforcement degrees of bottom, top and web reinforcement, respectively acknowledgements financial support from a grant-in-aid for scientific research from the japan society of promotion of science (19k15062) is greatly appreciated. references [1] r. c. elstner, e. hognestad, editors. laboratory investigation of rigid frame failure. journal proceedings, 53:637-668 1957. [2] s. b. bhide, m. p. collins. influence of axial tension on the shear capacity of reinforced concrete members. structural journal 86(5):570-81, 1989. [3] y. yang, j. den uijl, j. walraven. critical shear displacement theory: on the way to extending the scope of shear design and assessment for members without shear reinforcement. structural concrete 17(5):790-8, 2016. https://doi.org/10.1002/suco.201500135. [4] a. muttoni, m. fernández ruiz. shear strength of members without transverse reinforcement as function of critical shear crack width. aci structural journal 105(2):163-72, 2008. https://doi.org/10.14359/19731. [5] b. i. mihaylov, e. c. bentz, m. p. collins. two-parameter kinematic theory for shear behavior of deep beams. aci structural journal, 110(3):447-56, 2013. [6] fib. shear and punching shear in rc and frc elements (lausanne: international federation for structural concrete), fib bulletin 57:268, 2010. [7] sm. t. smith, d. a. howell, m. a. t. triska, et al. effects of axial tension on shear-moment capacity of full-scale reinforced concrete girders. aci structural journal 111(1):211, 2014. https://doi.org/10.14359/51686534. [8] d. fernández-montes, e. gonzález valle, e. díaz heredia. influence of axial tension on the shear strength of floor joists without transverse reinforcement. structural concrete 16(2):207-20, 2015. https://doi.org/10.1002/suco.201400063. [9] d. t. pham, b. fouré, n. pinoteau, et al. influence of axial tension on the shear strength of rc beams without stirrups. structural concrete 22(2):1037-53, 2020. https://doi.org/10.1002/suco.202000077. [10] k. h. tan, a. hasegawa, f. nishino. determination of ultimate shear strength of reinforced concrete beams without web reinforcement by a combined upper and lower bound analysis. doboku gakkai ronbunshu 1986(368):23-33, 1986. https://doi.org/10.2208/jscej.1986.368_23. [11] d. c. drucker. on structural concrete and the theorems of limit analysis. international association for bridge and structural engineering, division of engineering, brown university, p. 36, 1961. [12] s. j. chapman.matlab programming for engineers, cl engineering, toronto: nelson education, 2015. [13] aashto. aashto lrfd bridge design specifications, 7th ed. american association of state highway and transportation officials, washington, d.c., 2013. 289 https://doi.org/10.1002/suco.201500135 https://doi.org/10.14359/19731 https://doi.org/10.14359/51686534 https://doi.org/10.1002/suco.201400063 https://doi.org/10.1002/suco.202000077 https://doi.org/10.2208/jscej.1986.368_23 doi:10.14311/app.2021.30.0114 acta polytechnica ctu proceedings 30:114–120, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app application of the mori-tanaka method to describe the rate-dependent behavior of unidirectional fibrous composites soňa valentováa, michal šejnohaa, ∗, jan vorela, radek sedláčekb, pavel padevěta a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b czech technical university in prague, division of biomechanics, faculty of mechanical engineering, technická 4, 160 00 prague, czech republic ∗ corresponding author: sejnom@fsv.cvut.cz abstract. this paper examines the possibility of using the mori-tanaka micromechanical model describe the rate dependent behavior of the polymer matrix based fibrous composites. the generalized leonov model is adopted to capture the time and rate dependent character of the selected matrix, while fibers are assumed elastic. the performance of the mori-tanaka method is tested against the finite element simulations carried out in the framework of first-order homogenization. for simplicity, the periodic hexagonal array model is chosen to represent the microstructural arrangement of fibers in the yarn cross-section. to match the predictions provided by the two approaches a suitable modification to the original mori-tanaka method is proposed. an extensive parametric study is presented to illustrate a considerable improvement of the predictive capability of the modified mori-tanaka method. keywords: fibrous composites, generalized leonov model, homogenization, maxwell chain model, mori-tanaka method, periodic unit cell, viscoelasticity. 1. introduction unidirectional fiber-reinforced composites have already found a way into a variety of industrial fields. for such reasons as low weight, good thermal properties, high corrosion and chemical resistance, high longitudinal stiffness and strength, lower economic and environmental cost compared to other widely used materials, their application appears in aerospace, automotive, marine, and various construction sectors. in order to take full advantage of the benefits the composite material offers, a suitable computational model predicting its macroscopic response including the rate dependent behavior due to the viscous character of the matrix phase is necessary. despite a random nature of the distribution of fibers in the matrix, easily visible from images of material cross sections, a periodic hexagonal arrangement (pha) model [1] is adopted, because for high volume fraction of fibers it has been suggested as sufficiently accurate representation of a statistically homogeneous microstructure [2]. moreover, when limiting attention to elasticity the mori-tanaka (mt) method [3] delivers the macroscopic response comparable to finite element (fem) simulations assuming the pha model [4]. unfortunately, when moving beyond elasticity the original mori-tanaka scheme appears too stiff [5] and some action is needed to bring the predictions closer to those provided by fem [6, 7]. in the present study, we build upon the approach proposed in [8] for asphalts. however, their original formulations requires some modification to suit the present material system. while attention is primarily accorded to testing the mori-tanaka method as a very efficient substitute for a highly expensive finite element method in large scale simulation of textile composites [9], the material model representing the epoxy matrix is mentioned only briefly. further details on extensive experimental program to calibrate the leonov model [7, 8, 10] can be found in [11]. the paper is organized as follows. section 2 introduces the material system examined herein. next, section 3 shortly describes the leonov viscoelastic model adopted for the epoxy matrix. for the sake of convenience we briefly summarize the main results of the experimental program needed to calibrate the constitutive model. in particular, creep tests at different stress levels and rate dependent uniaxial tensile strain controlled experiments are described. finally, section 4 outlines two types of modifications of the mori-tanaka method accompanied by an extensive parametric study to examine their potential. the principal achievements are then summarized in section 5. 2. examined composite system following our recent study on the mori-tanaka method [5] we again consider a unidirectional fibrous 114 http://dx.doi.org/10.14311/app.2021.30.0114 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 viscoelastic modeling of unidirectional fibrous composites figure 1. cross section of examined fibrous composite. (a) (b) figure 2. a) mts alliance 30kn, b) polymer dog bone specimens. ea et ga gt νa cf [gpa] [gpa] [gpa] [gpa] [-] [-] carbon fibers 294 13 12 5 0.24 0.56 table 1. material properties of carbon fibers and its volume fraction in yarn. composite system made of elastic carbon fibers and nonlinear viscoelastic polymer matrix. a representative cross-section of the yarn and its binary form is displayed in the fig. 1. an image analysis was carried out to acquire the volume fraction of the fibers, see table 1 together with material properties taken from [9]. because fibers are transversely isotropic and the matrix is isotropic, the overall macroscopic response is expected to be also transversely isotropic. this is precisely what both the mt method and the pha model would predict in case of elastic response. while the mt method keeps this property even beyond elasticity, the pha model may yield the overall behavior orthotropic [5]. thus the application of the mt methods calls for caution when applied to real material systems. 3. generalized leonov model the generalized leonov model is adopted here to describe the nonlinear viscoelastic behavior of the epoxy matrix. the formulation assumes the bulk response be linearly elastic σm = kεv . (1) the viscoelastic response is thus driven solely by the deviatoric stress and strain components. for the maxwell chain model the corresponding constitutive equations read dsij dt = m! µ=1 2gµ( deij dt − de p,µ ij dt ), (2) sij = m! µ=1 s µ ij . (3) 115 s. valentová, m. šejnoha, j. vorel et al. acta polytechnica ctu proceedings figure 3. strain rate dependent uniaxial tensile tests. figure 4. eyring plot. the creep strain increment follows from the eyring flow model as dep dt = 1 2a sinh(τ /τ0) , (4) where a and τ0 are model parameters determined experimentally from either uniaxial [10, 12] or torsional tests [8] performed at different strain rates. for example, rewriting eq. (4) for a uniaxial tension we get σy = τ0 √ 3 ln(2a √ 3) + τ0 √ 3 ln ε, (5) which upon substituting the experimentally observed yield, or the maximum, tensile stress σy allows us to determine the necessary material parameters, see figures 3 and 4 and the next section. 3.1. strain rate dependent tensile tests six dog bone specimens in fig. 2(b) made of 285/500 aero havel epoxy resin were loaded in simple tension in the displacement controlled regime at a specific strain rate until failure using the mts alliance 30kn electromechanical testing machine equipped with 30kn load cell, see fig. 2(a). the resulting stress strain diagrams in fig. 3 confirm the rate dependent response of the epoxy matrix. replotting the yield stress as a function of the applied strain rate in a logarithmic scale and applying simple linear regression gives the material parameters of the eyring flow model in eq. (5). figure 5. creep experiments at various stress levels. figure 6. approximation of master curve. 3.2. creep tests the maxwell chain model requires calibration of the creep relaxation function typically found using the laplace transform of the creep compliance function. one-dimensional format of these functions can be written as j (t) = m! µ=1 jµ " 1 − exp # − t τµaσ (t) $% , (6) r(t) = m! µ=1 eµ exp # − t θµaσ (t) $ . (7) where τµ and θµ are the selected retardation and relaxation times, respectively, and aσ is the stress shift factor to define a nonlinear viscosity of a dashpot element in, e.g. the kelvin unit ηµ(τeq ) = τµjµaσ (τeq ), τeq = & 1/2sij sij . the compliances jµ can be derived by minimizing the least square difference of j (t) given by eq. (6) and experimentally constructed master curve for the selected set of retardation times, while exploiting the time-stress superposition principle. note that τ1 should be sufficiently small to represent in the limit τ → 0 a solid material. theoretically, τ1 = 0.001 might not satisfy this requirement but the resulting series validated in [11] proved satisfactory. to that end, a series of creep tests at different stress levels is performed first. the creep curves in fig. 5 are 116 vol. 30/2021 viscoelastic modeling of unidirectional fibrous composites (a) (b) figure 7. a) periodic unit cell (puc) with hexagonal arrangement of fibers in yarns, b) mesh of puc. µ τµ [s] jµ [mpa−1] θµ [mpa·s] eµ [mpa] 1 0.001 2.606512×10−4 9.927397×10−3 2.787166×101 2 0.01 1.905071×10−6 9.966502×10−2 1.278184×101 3 0.1 8.808431×10−7 9.815126×10−1 7.056602×101 4 1 4.934025×10−6 9.543319×10+0 1.711529×102 5 10 1.276165×10−5 9.344254×10+1 2.334448×101 6 100 1.969419×10−5 9.580883×10+2 1.418353×102 7 1000 1.290521×10−5 8.275395×10+3 5.659977×102 8 10000 6.291266×10−5 9.647045×10+4 1.586346×102 9 100000 7.887707×10−6 2.005373×10+5 1.944645×103 10 1000000 1.577867×10−3 4.168654×10+5 5.096147×102 table 2. parameters of maxwell chain model. then horizontally shifted to get the master curve for the selected reference stress. the experimental points as well as the resulting polynomial representation are plotted in fig. 6. ten maxwell units in parallel were considered in the present study to represent the maxwell chain model. the resulting parameters are listed in table 2. the shear moduli in eq. (2) are obtained from eµ for the selected poisson ratio set equal to 0.39. further details including validation of the proposed constitutive model are available in [11]. 4. homogenization as mentioned in the introductory part, we consider the predictions provided by fem as a benchmark to test the performance of the mt method. owing to a space limitation we present, with reference to fem simulations, the pha computational model only, see fig. 7, and refer the reader for details on theoretical grounds of the first-order homogenization to, e.g. [4, 13, and references therein]. 4.1. mori-tanaka method: standard formulation the mt method, however, deserves more attention. to begin with, we write the increments of local stress averages in individual phases as ∆σf = lf ∆εf , ∆σm = 'lm(∆εm − ∆µm), (8) where lr , r = f, m1, is the phase material stiffness matrix and 'lm represents the dependence on the current viscoelastic modulus. the local strain increments written in terms of the prescribed macroscopic strain increment ∆e and the increment of the creep strain ∆µm developed in the matrix follow from dvorak’s transformation field analysis, see [14], in the form ∆εf = 'af ∆e + 'df m∆µm, (9) ∆εm = 'am∆e + 'dmm∆µm. (10) where 'ar and 'drm are the mechanical strain localization factors and strain transformation influence functions, respectively. in the light of the mt method [3] they can be expressed as 'am = ( cmi + cf 'tf )−1 , 'af = 'tf 'am, (11) 'drm = * i − 'ar + * 'lm − lf +−1 'lm. (12) the partial strain concentration factor 'tf follows from the eshelby transformation field analysis and in general it is a function the shape and orientation of the fiber and instantaneous properties of the matrix. the principal drawback of the mt method is evident from fig. 8 comparing the fem simulations with 1subscripts f, m stand for the fiber and matrix phase, respectively. 117 s. valentová, m. šejnoha, j. vorel et al. acta polytechnica ctu proceedings figure 8. stress-strain diagrams for various uniaxial tensile strain rate loading. the mt results for various strain rates of the applied in-plane shear strain. it is reasonable to attribute the observed stiff response of the mt method to its inability to represent localization of inelastic strain since dealing with phase volume averages only. 4.2. mori-tanaka method: modified formulation owing to the previous results we propose a simple modification to match the predictions provided by both mt method and fem. being inspired by [8, 12] we attempt to describe the stress redistribution due to formation of shear bands by the reduction of stress taken by the fibers through a damage like parameter ω and write the fiber stress increment as ∆σf = (1 − ω)lf ∆εf . (13) two particular formulations describing the evolution of ω are investigated. the one suggested in [12] reads ω = 1 − " τ teq n τ0 /sinh( τ teq n τ0 ) %m , (14) where m, n are the model parameters and the equivalent stress τeq is given by τeq = , 1 2 sij sij . (15) the influence of parameters n and m is illustrated in fig. 9 suggesting a rather similar effect of the two parameters both in terms of stiffness and overall stress reduction. thus tuning only one of them appears sufficient in adjusting the macroscopic response. however, as seen in fig. 10, plotting the fem and mt estimates of the macroscopic response for the applied rate of macroscopic in-plane shear strain ėxy = 1 × 10−4 s−1 shows that such a formulation does not provide satisfactory results. therefore, we propose a new formulation written as ω = n " 1 − # τ teq − t τ0 /sinh( τ teq − t τ0 ) $%m , (16) (a) (b) figure 9. influence of parameters m, n in describing the evolution of ω according to eq. (14): (a) influence of parameter m , n =1, (b) influence of parameter n , m =1. figure 10. comparison of fem and mt predictions with ω given by eq. (14). where m, n, t are again the model parameters, which, unlike to eq. (14), provide more freedom in the model calibration. the way they control the evolution of ω is evident from fig. 11. while parameter m adjusts the rate of evolution of ω with τeq , the other two parameters n, t merely shift the damage curve horizontally and 118 vol. 30/2021 viscoelastic modeling of unidirectional fibrous composites (a) (b) (c) figure 11. influence of parameters m, n in describing the evolution of ω according to eq. (16): (a) influence of parameter m , n =1, t =0; (b) influence of parameter n , m =1, t =0; (c) influence of parameter t , m =n =1. vertically. thus n controls the minimum stress the fibers may take and t serves to delay the onset of damage evolution. note that these two parameters are subject to some limitations, particularly n = 〈0, 1〉, (17) and if τeq < t then ω = 0, (18) figure 12. comparison of fem and mt predictions with ω given by eq. (16). if τeq > t then ω follows from eq. (16). (19) figure 12, showing again the macroscopic stressstrain curve for the applied shear strain rate ėxy = 1 × 10−4, promotes this new formulation as almost a perfect match between fem and mt predictions can be obtained with properly adjusted parameters m, n, t in eq. (16). 5. conclusions the main goal of this study was to confirm suitability of the mori-tanaka method to replace demanding finite element simulations in the prediction of the macroscopic response of fibrous composites. this would prove advantageous particularly in a multiscale analysis of textile composites where the response at the level of yarn must be solved independently from the textile ply. to that end, two modifications to the original mt formulation were tested with particular attention devoted to the nonlinear viscoelastic response of the matrix phase governed by the generalized leonov model. both modifications attempt to reduce the stress taken by the fibers by introducing a damage like parameter ω in the constitutive equation for the fiber phase. the first modification originated from the work of valenta et al. [8, 12]. while successful for asphalt mixtures it proved insufficient in the present study. therefore, a new modification was proposed to allow for more freedom in the calibration of the evolution law for damage parameter ω. the results clearly show a superiority of this new formulation allowing us to receive a very good agreement between fem and mt predictions. this also promotes the mt method as a perfect tool for the prediction of macroscopic response of two-phase composites made of aligned fibers embedded into a polymer matrix. acknowledgements the financial support provided by the gačr grant no. 19-15666s and by the czech technical university in 119 s. valentová, m. šejnoha, j. vorel et al. acta polytechnica ctu proceedings prague within sgs project with the application registered under the no. sgs20/038/ohk1/1t/11 is gratefully acknowledged. references [1] j. l. teplý, g. j. dvorak. bound on overall instantaneous properties of elastic-plastic composites. journal of the mechanics and physics of solids 36(1):29–58, 1988. [2] j. zeman, m. šejnoha. numerical evaluation of effective properties of graphite fiber tow impregnated by polymer matrix. journal of the mechanics and physics of solids 49(1):69–90, 2001. [3] y. benveniste. a new approach to the application of mori-tanaka theory in composite materials. mechanics of materials 6:147–157, 1987. [4] m. šejnoha, j. zeman. micromechanics in practice. wit press, southampton, boston, 2013. [5] s. valentová, m. šejnoha, j. vorel. comparing moritanaka method and first-order homogenization scheme in the viscoelastic modeling of unidirectional fibrous composites. acta polytechnica ctu proceedings 26:133–138, 2020. doi:10.14311/app.2020.26.0133. [6] r. masson, m. bornert, p. suquet, a. zaoui. an affine formulation for the prediction of the effective properties of nonlinear composites and polycrystals. journal of the mechanics and physics of solids 48:1203–1227, 2000. [7] m. šejnoha, r. valenta, j. zeman. nonlinear viscoelastic analysis of statistically homogeneous random composites. international journal for multiscale computational engineering 2(4):645–673, 2004. [8] r. valenta, m. šejnoha, j. zeman. macroscopic constitutive law for mastic asphalt mixtures from multiscale modeling. international journal for multiscale computational engineering 8(1):131–149, 2010. [9] j. vorel, e. grippon, m. šejnoha. effective thermoelastic properties of polysiloxane matrix based plain weave textile composites. international journal for multiscale computational engineering 13(3):181–200, 2015. [10] t. a. tervoort. constitutive modeling of polymer glasses: finite, nonlinear visocelastic behaviour of polycarbonate. ph.d. thesis, eindhoven university of technology, eindhoven, 1996. [11] m. šejnoha, s. valentová, j. vorel, et al. modeling of viscoelastic response of unidirectional fibrous composites made of basalt and carbon fibers. in w. p. de wilde, s. hernandes (eds.), 10th international conference on high performance and optimum design of structures and materials. wit press elibrary, https://www.witpress.com/elibrary, 2020. [12] r. valenta, m. šejnoha. hierarchical modeling of mastic asphalt in layered road structures based on the mori-tanaka method. acta polytechnica 52(6):48–58, 2012. [13] s. valentová, v. hrbek, j. vorel, m. šejnoha. strength of composite yarn under biaxial loading. acta polytechnica ctu proceedings 15:131–136, 2018. [14] g. j. dvorak, y. benveniste. on transformation strains and uniform fields in multiphase elastic media. proceedings of the royal society of london series a mathematical, physical and engineering sciences 437(1900):291–310, 1992. 120 https://doi.org/10.14311/app.2020.26.0133 doi:10.14311/app.2021.30.0047 acta polytechnica ctu proceedings 30:47–52, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app localization analysis of damage for one-dimensional peridynamic model karel mikeša, ∗, milan jiráseka, jan zemana, ondřej rokošb, ron h. j. peerlingsb a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague, czech republic b eindhoven university of technology, department of mechanical engineering, po box 513, 5600 mb eindhoven, the netherlands ∗ corresponding author: karel.mikes@cvut.cz abstract. peridynamics is a recently developed extension of continuum mechanics, which replaces the traditional concept of stress by force interactions between material points at a finite distance. the peridynamic continuum is thus intrinsically nonlocal. in this contribution, a bond-based peridynamic model with elastic-brittle interactions is considered and the critical strain is defined for each bond as a function of its length. various forms of length functions are employed to achieve a variety of macroscopic responses. a detailed study of three different localization mechanisms is performed for a one-dimensional periodic unit cell. furthermore, a convergence study of the adopted finite element discretization of the peridynamic model is provided and an effective event-driven numerical algorithm is described. keywords: damage, localization, nonlocal continuum, peridynamics. 1. bond-based peridynamics for natural as well as man-made materials, the resulting macroscopic response is dictated by the finescale structure of the material. individual building units (such as atoms on the nanoscale, inclusions and matrix on the microscale, or bricks on the mesoscale) have a certain characteristic size and their interaction takes place at finite distance. this can be taken into account by particle/lattice models, or by various enriched continuum theories. in this work, we consider the bond-based version of the peridynamic model and study the effect of the specific choice of the critical strain function on the evolution and localization of the macroscopic damage. peridynamics was introduced in 2000 by s. a. silling as a reformulation of the classical theory of elasticity [1]. in peridynamics, the partial differential equations of the classical continuum theory are replaced by integral equations. since differentiation of the displacement field is not needed, peridynamics is especially convenient for capturing discontinuities such as cracks without the complications of mathematical singularities. in a peridynamic model, the material is represented by infinitely many infinitely small particles that can interact at a finite distance. usually, interaction between two points is considered only if their distance does not exceed the so-called horizon h, see figure 1. the acceleration ü of a particle located at a spatial position x in the reference configuration at time t can h figure 1. two-dimensional representation of a discretized peridynamic model: regular grid of particles (black dots) and one selected particle (hollow circle) with interactions (grey lines) within horizon h (dashed line). be found from the peridynamic equation of motion ρü(x, t) = ! hx f (u(x′, t) − u(x, t), x′ − x)dvx′ + b(x, t) (1) ρü(x, t) = ! hx f (u(x′, t)−u(x, t), x′−x)dvx′ + b(x, t) where hx is the set of points that interact with the particle at x, u is the displacement vector field, b 47 http://dx.doi.org/10.14311/app.2021.30.0047 http://ojs.cvut.cz/ojs/index.php/app k. mikeš, m. jirásek, j. zeman et al. acta polytechnica ctu proceedings ⇠ h�h s0(⇠) ⇠ h�h s0(⇠) ⇠ h�h s0(⇠) figure 2. various functions describing the dependence of failure strain on the initial bond length: constant (left), decreasing (middle), and increasing (right). is the prescribed body force density field, ρ is the mass density in the reference configuration, and f is the so-called pairwise force function. the value of f is the force vector (per unit volume squared) that a particle located at the position x′ exerts on the particle located at x. the pairwise force function can be written in the form f (η, ξ) = ξ + η |ξ + η| f (|ξ + η|, ξ) (2) where ξ = x′ − x (3) is the relative position of the interacting particles in the reference configuration, η = u(x′, t) − u(x, t) (4) is their relative displacement, and f is the scalar bond force governed by a constitutive relation. in this work, bonds are considered as perfectly elastic-brittle, in line with the concept of microelastic brittle material, adopted from [2]. the corresponding bond force is defined as f (|ξ + η|, ξ) = c s(|ξ + η|, |ξ|) µ(t, ξ) (5) where c is a spring constant called the bond elastic micromodulus, s(|ξ + η|, |ξ|) = |ξ + η| − |ξ| |ξ| (6) is the bond stretch and µ(t, ξ) = " 1 if s(t′, ξ) < s0 for all 0 ≤ t′ ≤ t, 0 otherwise (7) is a history-dependent damage function that jumps from 1 to 0 when the bond breaks, which is assumed to happen when its stretch attains the critical level, s0. in this work, a fixed value of the elastic micromodulus c is considered for all bonds, whereas the critical stretch for bond failure, s0, is evaluated for each link as a function of its length. three different forms of length functions, namely constant, linear decreasing, and linear increasing, are employed, as illustrated in figure 2. l ux h = 0.3lx 0.1l ⇠ = 0.1l ⇠ = 0.2l ⇠ = 0.3l figure 3. geometry of an illustrative example of the one-dimensional finite element model of a periodic unit cell of length l with horizon h = 0.3l: primary nodes (filled circles), slave nodes (hollow circles), virtual node with prescribed displacement (gray circle), and an expanded illustration of the considered interactions with different lengths ξ (gray solid lines). dashed arrows denote connections of slave nodes with primary nodes. 2. numerical implementation for numerical implementation, we use an analogy with the finite element analysis, adopted from [3], to represent a bond-based peridynamic system by a finite element truss model. thus, individual bonds on the microscopic level are modelled by truss elements, so-called links, and the material parameters of the bonds, c and s0, are related to the macroscopic young modulus e and critical strain ε0, respectively. 2.1. one-dimensional model damage localization and evolution of failure mechanisms in a one-dimensional periodic lattice submitted to positive strain is investigated for different forms of the critical strain length function. the geometry of an illustrative example of the finite element model of a one-dimensional periodic unit cell is depicted in figure 3. all nodes in the periodic unit cell (except the last one) are modeled as primary nodes that carry independent degrees of freedom (horizontal displacements). to enforce periodicity, additional slave nodes are constructed as periodic images of primary nodes on one side of the periodic cell. the number of periodic slave nodes is given by the size of the horizon. loading in the form of prescribed displacement is applied to the cell through one virtual node which is linked to all slave nodes simultaneously. the displacement of each slave node is evaluated as the sum of the unknown displacement of its master (one of the primary nodes) and the prescribed displacement of the virtual node. bonds, in the form of one-dimensional truss elements, are placed between all pairs of primary nodes as well as between primary nodes and slave nodes within the given horizon h; see the solid grey lines in figure 3. the cross-sectional area of a truss element is the same for all links and it is scaled to provide a unit sum of overlapping link areas representing the macroscopic cross-sectional area. 48 vol. 30/2021 one-dimensional peridynamic model damage analysis during the loading process, the evolution of damage in a macroscopic cell is modelled as a sequence of failures (breakages) of individual links followed by stress redistribution. the macroscopic damage ω can be evaluated for each point x of the unit cell as ω(x) = nf (x) n (x) (8) where n is the number of all links passing through the given point x and nf is the number of broken links (i.e., those with µ = 0) passing through x. theoretically, it is possible that multiple links can reach the critical strain simultaneously. then the subsequent solution is not unique and additional choices must be made to distinguish one link or a group of links preferred for failure in the given step. for an ideal lattice geometry without imperfections, such a situation occurs quite often, not only at the beginning of the loading, but also almost every time when broken links form a symmetric pattern. to avoid this phenomenon and to more closely mimic the behavior of real materials, the initial position of the primary nodes is randomized by a small perturbation with a uniform distribution and a maximum value l × 10−6. the macroscopic response of the randomized model is evaluated statistically based on multiple random realizations. 2.2. solution algorithm to capture the exact sequence of breaking links, it is important to investigate not only the beginning of failure but also the full failure process of individual links and take into account that a new link can start failing before another (already failing) link becomes fully stress-free. for elastic-brittle links, failure means breaking, i.e., decreasing link stress under constant link strain. in this case, neglecting the possibility of partial unloading usually has only a minor effect on the macroscopic response. however, for elastic-damaging links with softening (which are not considered in this contribution), the possibility of partial unloading is essential. it is obvious that the size of the loading increment to reach link failure or unloading is different in each step. therefore, conventional solvers for an incremental nonlinear structural analysis are inconvenient for the present problem because they may miss the correct failure mechanism if the prescribed increment is too large. to solve the present problem exactly, an event-driven algorithm with step size control introduced in [4] has been employed. since the bond-level constitutive law is piecewise linear, the step size can be selected such that the tangential stiffness matrix does not change during the step. therefore, no equilibrium iteration is necessary. in one solution step, either a positive or a negative force increment is considered to achieve the onset of failure or complete failure of one link. furthermore, an effective method of inelastic forces can be used to avoid the necessity of decomposing the tangential stiffness matrix in each step; for more details see section 4.2 in [4]. 3. results in this section, results are presented for the aforementioned finite element representation of the peridynamic model of a one-dimensional periodic cell with elastic-brittle interactions subjected to positive (average) strain. the modelled cell is considered to have a length of l = 1 m, with the horizon h = 0.1l and cross-sectional area a = l2. the links have a constant young modulus e = 100 n m−2 and critical strain function with a maximum value ε0 = 1 × 10−2 and various forms, namely constant, linearly decreasing, and linearly increasing with link length. 3.1. convergence study first of all, a convergence study of the peridynamic model with decreasing computational grid spacing is performed. for that purpose, five simulations with the same macroscopic parameters but with 10, 20, 50, 100, and 200 primary nodes have been performed. the simplest case of constant critical strain for different link lengths has been considered. the results are shown in figure 4 in terms of normalized forcedisplacement diagrams, in which the elastic part of the displacement, f/kel, is subtracted from the total macroscopic displacement u and only the inelastic part is plotted. f is the macroscopic reaction force and kel denotes the initial elastic stiffness. for the purpose of clarity, only one representative random realization is shown for each discretization. figure 4. convergence of normalized forcedisplacement diagrams (elastic part of displacements has been subtracted) for one-dimensional periodic lattice with fixed size l, horizon h = 0.1l, for discretizations with different numbers of nodes. for each discretization, only one representative random realization is shown. the maximum force corresponding to the initial failure of the first link as well as the maximum displacement for the final damage mechanism are predicted exactly for all refinements. however, for a small number of nodes, significant oscillation can be 49 k. mikeš, m. jirásek, j. zeman et al. acta polytechnica ctu proceedings figure 5. results for constant critical strain: final damage profile (top left), positions and lengths of broken links (bottom left), and normalized force-displacement diagram where the elastic part of the displacement has been subtracted (right). figure 6. results for critical strain linearly decreasing with link length: final damage profile (top left), positions and lengths of broken links (bottom left), and normalized force-displacement diagram where the elastic part of the displacement has been subtracted (right). observed, while with an increasing number of nodes, the macroscopic response tends to a smooth curve. discretization with 100 nodes turned out to be sufficient to capture the important characteristics of the macroscopic response. note that for 10 primary nodes, there is only one link between each pair of neighbouring nodes and therefore damage localizes in one step by failure of the first link and the corresponding response is a straight line, see the gray line in figure 4. 3.2. damage mechanisms in this subsection, the influence of the form of the critical strain function is analyzed. the discretization with 100 nodes is used based on the previous study. the results are shown in figures 5, 6, and 7 for the constant, decreasing, and increasing functions of the critical strain, respectively. for each function, 10 independent random realizations have been evaluated to analyze the influence of randomness to the mechanisms of failure. however, for the purpose of clarity, only the results of one representative random realization are depicted. the resulting mechanisms of failure are illustrated by the final damage profiles (top left in each figure), evaluated according to eq. (8), and by the positions of all broken links (bottom left); here, each blue dot represent one broken link. the length of the broken link ξ/l is plotted against the position of its center. the values are normalized by the cell size l. moreover, the corresponding inelastic force-displacement diagram (right) is shown to illustrate the macroscopic 50 vol. 30/2021 one-dimensional peridynamic model damage analysis figure 7. results for critical strain linearly increasing with link length: final damage profile (top left), positions and lengths of broken links (bottom left), and normalized force-displacement diagram where the elastic part of the displacement has been subtracted (right). behavior during damage evolution. 3.2.1. constant critical strain in the initial (undamaged) configuration, the strain of the links is uniform. consequently, for the first case of constant critical strain, there is no preferred length of links for failure in the first step. in our simulations, one of the shortest links is made weaker to initialize damage localization. once the shortest link breaks, the stress is redistributed, two links in the second shortest layer spanning the broken link become the most loaded links and one of them fails. then, each subsequent broken link is situated in the next longer layer spanning all previously broken links until the first link in the longest layer is reached. once this sequence of propagating broken links from the shortest to the longest layer is finished, similar sequences follow starting from the second shortest to the longest, the third shortest to the longest, etc., until the final triangular pattern visible in figure 5 (bottom left) is formed. note that if the shortest layer of links is not artificially favored at the beginning, the initial failure can occur for a different link length due to the randomization. however, in such a case, the closest shortest links break in the next step and the same failure sequences can be observed with only the exception of the first link. for constant critical strain, the position of maximum damage (ω = 1), i.e., the position of macroscopic failure, is dictated by the first broken link. the final damage mechanism is simple. the resulting failure pattern and associated failure properties such as fracture energy can be evaluated from the geometry of the cell. also the force-displacement diagram does not exhibit any significant irregularities. 3.2.2. decreasing critical strain in the case of a critical strain decreasing with link length, the longest layer of links is the weakest one and the first broken link appears in this layer. subsequently, other links in the weakest layer are breaking one by one until the weakest layer is completely broken. this process is repeated for each next longest layer of links until a critical localization layer is reached. the gradual breaking of individual layers corresponds to horizontal parts of individual steps (jumps) in the force-displacement diagram. particularly, four steps can be observed at the levels of normalized force about 0.01×10−2, 0.08×10−2, 0.12×10−2, and 0.135×10−2, see fig. 6 (right). from the macroscopic point of view, this process can be understood as a uniform growth of damage. once the critical localization layer is reached, broken links localize in a triangular pattern similar to the previous case of constant critical strain; see fig. 6 (bottom left). in the force-displacement diagram, localization corresponds to the final unloading part, where both macroscopic force and macroscopic displacement are decreasing. the position of maximum damage is not affected by the first broken link but it is determined during the localization of the damage in the critical localization layer. 3.2.3. increasing critical strain for the last case of increasing critical strain, the shortest layer of links is the weakest one and the first broken link appears here. however, in this case, damage of the first link is followed neither by damage of the whole weakest layer nor by localization. on the contrary, links continue breaking in different layers without a specific pattern. consequently, the macroscopic damage is distributed randomly and no clear 51 k. mikeš, m. jirásek, j. zeman et al. acta polytechnica ctu proceedings localization is observed. in the end, the majority of broken links are situated in several weakest layers. however, these weak layers (except the last one) are not completely broken. on the other hand, in several regions, damage significantly propagates also into layers of stronger links, see fig. 7 (bottom left). in the central region, the broken links form an unusual mechanism of failure where no macroscopic cross-section exhibits complete damage (ω = 1), even when the macroscopic reaction force completely vanishes. such a mechanism is formed by overlapping sliding parts. as a result of this, the final macroscopic strain is localized not only into one but into three segments between neighbouring nodes where two of them exhibit positive strain increments and the remaining one a negative strain increment of equal magnitude. unlike the previous cases of constant and decreasing forms of critical strain function, in this case, the order of broken links is unstable and the position and shape of the final damage pattern as well as the evolution of the damage before localization are affected by the randomization of the initial lattice geometry. as a result, significant oscillations can be observed in the force-displacement diagram, see fig. 7 (right). 4. conclusions in this contribution, a finite element truss model representing a peridynamic system with elastic-brittle interactions has been introduced, together with an accurate and efficient event-driven numerical algorithm which is able to capture exactly not only the failure but also the full unloading of individual links. the convergence study of the presented model has shown that upon refinement the system tends to a smooth macroscopic response. the influence of various forms of the critical strain function on the type of macroscopic response has been investigated for a one-dimensional periodic unit cell. constant, decreasing, and increasing critical strain functions have been investigated and the resulting responses with different types of damage evolution have been analyzed, namely direct damage localization, uniform damage evolution preceding localization, and random damage distribution, respectively. the presented results demonstrate that the concept of critical strain values depending on link length can be used to reach various types of macroscopic responses even for such a simple geometry (1d) and one of the simplest nonlinear material models (elasticbrittle). the significant advantage of this concept is that the presented elastic-brittle model can be fitted to represent various macroscopic behaviors, but at the same time the micro-level constitutive law remains piecewise linear and therefore the exact and efficient event-driven algorithm can be used. in addition, the spectrum of different response types can be further extended by considering various micromodulus functions or by adopting elasticdamaging links with softening. list of symbols h material horizon [m] hx neighborhood of a particle vx volume of a particle neighborhood x particle coordinates [m] u displacement vector [m] ü acceleration vector [m s−2] ρ mass density [kg m−3] b body force density [n m−3] t time [s] ξ relative position of two particles [m] η relative displacement of two particles [m] f pairwise force function [n m−6] f scalar bond force [n m−6] s bond stretch [–] x horizontal coordinate [m] l cell size [m] a cross-sectional area of links [m2] ux prescribed horizontal displacement [m] u total lattice displacement [m] f reaction force [n] kel initial elastic stiffness [n m−1] c spring constant [n m−2] s0 critical stretch for bond failure [–] e young modulus [n m−2] ε0 critical strain [–] µ damage function [–] ω damage [–] n number of links [–] nf number of broken links [–] acknowledgements this research has been performed in the center of advanced applied sciences (caas), financially supported by the european regional development fund (project no. cz.02.1.01/0.0/0.0/16_019/0000778). financial support received from the czech technical university in prague under project sgs20/038/ohk1/1t/11 is gratefully acknowledged. references [1] s. a. silling. reformulation of elasticity theory for discontinuities and long-range forces. journal of the mechanics and physics of solids 48(1):175–209, 2000. [2] s. a. silling, e. askari. a meshfree method based on the peridynamic model of solid mechanics. computers & structures 83(17-18):1526–1535, 2005. [3] r. w. macek, s. a. silling. peridynamics via finite element analysis. finite elements in analysis and design 43(15):1169–1178, 2007. [4] m. jirásek, z. p. bažant. macroscopic fracture characteristics of random particle systems. international journal of fracture 69(3):201–228, 1994. 52 acta polytechnica doi:10.14311/app.2020.27.0042 acta polytechnica 27(0):42–47, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app micro-scale study of residual stresses in cr2o3 coatings sprayed by aps franck decroosa, ∗, cécile langladea, eric bourillota, geoffrey daruta, manuel francoisb a université de bourgogne franche-comté, laboratoire interdisciplinaire carnot de bourgogne, site de sevenans, 90 010 belfort cedex, france b université de technologie de troyes, charles delaunay institute, life assement of structures, materials and integrated systems, 12 rue marie curie, 10 004 troyes cedex, france ∗ corresponding author: franck.decroos@utbm.fr abstract. whichever the application field, every material forming process generates residual stresses on the surface. while they are likely to enhance the aimed properties of the final mechanical part, these stresses may also drastically reduce them and result in early failures. therefore, understanding the residual stress state remains a major challenge when coating complex parts, especially as most characterization methods at the microscopic scale involve specific sample preparation procedures which may affect the residual stresses field. this work investigates the residual stress state that exists in chromium oxide coatings deposited via atmospheric plasma spray (aps), using two pioneering techniques featuring high spatial resolution: scanning microwave microscopy and raman microspectroscopy. the first technique combines the measurement of microwave electromagnetic capacities of a vector network analyzer with the subnanometric resolution of an atomic force microscope: it thus enables performing depth investigations at very accurately defined positions of the probe on the surface. the second technique relies on the principle of photons inelastic scattering and involves a laser beam aiming at the material sample: it allows a fine characterization of the microstructure as well as defects and stresses detection via molecular vibratory signatures. the investigation scale is limited here to a few cubic micrometers. due to the highly localized scales of our investigations, which also depend on the device, the objective of our procedure required that the comparison should be made not on individual points but on definite mapped areas, every spot being analyzed and post-treated one after another, with optimum measuring parameters. results have been correlated with xrd measurements to cross-check the average amount of stress observed over a wider area. keywords: coatings, chromium oxides, raman micro-spectroscopy, residual stress, scanning microwave microscopy. 1. introduction 1.1. residual stresses in thermal spray coatings it is today widely-known that residual stresses (rs) have significant implications in both the preparation and the performance of thermally sprayed coatings. on one hand, tensile stresses are generally recognized to reduce service lifetime, weaken interfacial adhesion, and sometimes cause early delamination when they exceed the elastic limit of the material. on the other hand, compressive stresses are most often associated with better performance and fewer micro-cracks, which are known to enhance corrosion resistance during service [1–3]. thus, assessing the rs field of a coating is a major concern to optimize spraying parameters and match technical specifications. by doing so, a more accurate prediction of the in-service behavior of the coating is possible. important properties of rs are magnitude, sign, and even more specifically their superficial distribution. 1.2. origins of residual stresses thermal spray processes involve various stress generating phenomena. when dealing with ceramic coatings deposited via atmospheric plasma spray (aps), the most significant are the following. the first one is the effect of quenching, when the shrinkage of individual splats leads to local stress generation as the coating is cooling from its melting temperature to the substrate temperature. simultaneously, stressrelaxation mechanisms such as interfacial sliding and extensive microcracking may occur, microcracking being very common in ceramic materials [4–7]. the second phenomenon is due to the mismatch between the coefficients of thermal expansion (cte) of the coating and the substrate [1], [8]. a further effect lies in the transient injection of heat fluxes into the substrate, hence generating important thermal gradients which cause differential thermal expansion at different depths as well as changes to the local misfit strains [1]. this point is important in our study as the depth of investigation varies for the two tech42 http://dx.doi.org/10.14311/app.2020.27.0042 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 residual stresses in cr2o3 coatings niques employed. macro stresses (type 1 rs) and joint stresses (type 2 rs) are worth investigating in our case in consideration of the engineering objective [9]. 1.3. purpose of the work in practice, the superimposition of the various phenomena mentioned above generates a complex rs field and makes the in-service behavior of aps coatings difficult to predict. therefore, the objective of this work is to combine new investigation methods for local determination of the rs field, targeting the two kinds of stress mentioned above. the first method is the raman micro-spectroscopy, which is increasingly used in this field [9–14]. this technology enables fast micro-scale measurements on raman-active materials such as oxides [9, 10].here, lattice bondings will be submitted to vibration through the inelastic photon scattering phenomenon using a low power laser emission. hence, all molecules are characterized by a unique spectrum. the fingerprint spectrum and behavior law of raman peaks regarding stress are today known for common types of materials including cr2o3 [11, 12]. the second method is a custom scanning microwave microscope (smm) performing nano or micro-scale measurements, here used for the first time on ceramics [13]. hence, the study also aims to calibrate this new instrument using raman spectrometer as reference. smm features two advantages: a 20 nm diameter area of investigation at surface, and the data collection at different depths underneath this small area. moreover, this is a non-destructive investigation method. indeed, rs measurement is assessed while measuring the signal amplitude, which will vary with changes in lattice organization of the investigated material. a more complete theory is available in “experimental section” of aybeke, en and al. paper [14]. this in-depth technique represents a significant change in stress investigation, especially when dealing with ceramic coatings. finally, x-ray diffraction (xrd) measurements based on the sin2(ψ) method will be correlated to the previous results. xrd stands for the reference in our field [1, 9]. to do so, mappings have been performed with both more local devices in order to be compared to xrd wider area. raman spectrometer and top surface smm mappings results can be expected to exhibit similar tendencies as both investigate changes in lattice organization. these investigation technologies have been applied on cr2o3 coatings performed by aps: the spraying temperature was chosen as the varying parameter to elaborate different residual stress fields. 2. methods 2.1. materials and coating deposition chromium oxide coatings were deposited via aps upon a 30 mm × 30 mm × 5 mm steel substrate using commercially available chromium oxide powder from saint-gobain. before deposition, substrates were sandblasted so that the ra was superior to 4 µm, then they were cleaned in ethanol. in order to investigate differences between rs fields, this paper deals with two samples: (1.) ref1 which was performed using optimized parameters and procedure for thermal barriers coating applications at a depositing temperature of 200 ◦c. thickness is 200 µm. (2.) ht2 which only differs from the previous one by its depositing temperature that was about 450 ◦c. thickness is 400 µm. the main spraying parameters are listed in table 1. measurements were conducted on as-deposited coatings. spraying parameters value current intensity (a) 650 ar/h2 ratio 40/13 gas flow rate (l.min-1) 3 spraying distance (mm) 110 table 1. main spraying parameters 2.2. scanning microwave microscopy the used scanning microwave microscope is an original device from agilent which has been widely customized since it was bought. it is basically made out of two parts as showed in figure 1. figure 1. scanning microwave microscope the first one is an atomic force microscope (afm), equipped with a probe which will be used as an antenna to emit as well as to receive the microwave signal. to do so, being in the near field becomes a necessary condition in order to emit a directive signal. thus, the tip of the probe was put directly in contact with the sample surface. besides, the design of the tip also guarantees the high resolution of the measurement as its diameter is no more than 20 nanometers. the second part of the device is a vector network analyzer (vna). at first, this vna generates a wave train of a sinusoidal microwave signal at a very low power. then, it compares the generated signal to the one which is reflected. in short, vna outputs the s11 reflection coefficient. as a matter of fact, generating a wave train of various frequencies enables in-depth investigation of 43 f. decroos, c. langlade, e. bourillot et al. acta polytechnica figure 2. a typical microwave signal a material. hence, a finite number of depths can be investigated, all of them corresponding to specific resonant frequencies which are machine dependent. the lower the frequency, the deeper the investigation underneath the surface. a typical spectrum of scanning microwave microscopy is shown as an example on figure 2 with its resonant frequencies denoted by asterisks. last but not least, it is commonly acknowledged that stress variation inside a material will lead to a variation in amplitude of the microwave signal [12, 13]. in brief, this is exactly the ordinate variation of the asterisks that are studied in this paper. in practice, measurements were carried out from 0.5 ghz to 13 ghz, which means that the waves go from the surface coating (13 ghz) down to the substrate (0.8 ghz). the temperature was kept between 22.3◦c to 24◦c and the air humidity between 25 % to 27 %. 2.3. raman micro-spectroscopy the raman micro-spectroscopy examinations were performed with a xplora plus from horiba which was calibrated with a (111) si single crystal. measurements were performed using the parameters listed in table 2. the diameter of the irradiating laser spot is about 0.87 µm and depth penetration is about 0.380 µm. laser irradiance was kept under 1 gw as advised in previous works [14]. parameters value acquisition time (s) 15 accumulations 15 objective x50 gratings 1800 filter (%) 10 table 2. raman spectroscopy parameters this paper focuses on the shift of the most intense mode 555 cm−1 a1g vibration [15, 16] as it is recogfigure 3. edge effects in smm nized to produce a large change in the effective mode volume, thus undergoing the largest stress-induced frequency shifts [17]. the peak was fitted using either the pseudo-voigt or the lorentzian curve model based on the variation of the cumulative residuals. iterative curve fitting (icf) was performed under origin2016. 2.4. x-ray diffraction measurements were performed with a d8 discover from bruker equipped with a cu source, in ψ (or χ) type assembly configuration and θ-2θ set up. the post treatment was carried out with commercial software leptos using the sin2 ψ method. we focused on the shift of the peak at 125◦ with a peak evaluation chosen as standard and a stress model defined as normal in the software. 3. results and discussion 3.1. x-ray diffraction results at first, xrd measurements highlighted that there were very low amounts of stress at the surface of both samples, although the one coated at high temperature seems to be in a slightly more compressive state as shown in table 3: sample stress value(mpa) measurement uncertainty (mpa) ref1 5.0 ± 10.9 ht2 -52.4 ± 4.5 table 3. x-ray diffraction results 4. scanning microwave microscopy results 4.1. stress variation over centerlines an exploration along the centerline involving about thirty measurements was performed on both samples. 44 vol. 27/2020 residual stresses in cr2o3 coatings figure 4. averaged amplitude of smm signal against frequency results show that edge effects like the one visible on figure 3 at 9.1 ghz are likely to be found between 7.1 ghz and 10.9 ghz. change in solidification process may indeed occur near the edges of the sample as the number of degrees of freedom increases. this phenomenon is also to be considered when investigating closer to the very last sprayed layer where solidification was left to happen freely at ambient temperature. steel substrate is likely to exhibit edge effects too regarding cutting and drilling processes it underwent. in order to compare the state of stress at different depths between the two samples in a sound and relevant manner, only data between -5 mm to +5 mm were considered for the in-depth stress study. 5. in-depth stress variation we computed the average amplitude value obtained for each frequency within the delimited area. we therefore plotted the average stress value against the working frequency which reflects the investigation depth in the material. results are shown in figure 4: one can easily notice that both samples can be considered isotropic as levels of amplitude variate the same along the two directions. this result is consistent with what is expected from aps. moreover, the four curves exhibit the same steadily stress level as we investigate the [8-0.8] ghz range. such a tendency leads to think that this reflected signal characterizes the substrate, which is consistent with results from deep hole drilling investigation on thermally sprayed ceramics [2]. around the 9 ghz region would be the interface, whereas the 11.9 ghz to 10.9 ghz characterize the stress level within chromium oxides. collecting such few data in the coating can be explained by the evanescent behavior of microwave within ceramics. in concrete terms, the wave train easily penetrates the material following an exponential law of propagation as mentioned in literature [18]. this could also explain that, despite the thickness difference between ref1 and ht2, we collected the same amount of data in the two coatings. the authors want to set clear that smm resonant frequencies are very unlikely to match with the same depth of investigation between the two samples. 6. raman spectroscopy results 6.1. stress variations over centerlines centerlines stress measurements were performed with raman spectrometer to be compared to highest frequency (11.9 ghz) smm results. the resulting raw raman data are shown in dashed line on figure 5. 45 f. decroos, c. langlade, e. bourillot et al. acta polytechnica figure 5. correlation of stress measurements along y centerlines for the two samples to highlight clearer tendencies from them, a 5-points smoothing was applied, which does not impact curve extremities. from ht2 y, raman data reveal a strong wavy curve especially in the 6 millimeters area from either side of the center which is also exhibited by smm results. regarding ref1 y, raw raman data oscillate around 551.25 cm−1 from -12 mm to 0 mm, while smm data does around -12.75 db. then, as raman data decrease down to 548.5 cm−1, smm ones increase up to -12.3 db. at the last extremity of the curve, raman data drastically increase whereas smm ones start to decrease. although in practice stage position between the two instruments cannot perfectly match, resulting stress curves exhibit strong tendencies to face each other like mirror images. as the zero-stress reference could not have been reliably determined yet, authors find it too hasty to establish the borderline between compressive and tensile stresses in this paper. 7. conclusions the present work establishes that smm and raman spectrometer reveal similar stress tendencies at the center of thermally sprayed ceramic coatings. the difference between the volume investigated by each two instruments may explain the dissimilar results closer to the edge. as the smm showed consistent results regarding both top surface and in-depth investigations, a further study will assess microwave propagation law within cr2o3. acknowledgements this work is performed under the support of région bourgogne franche-comté under the research agreement number 2017y -06236/06807. the raman microscopy equipment is part of the mifhysto platform also supported by the region bourgogne franche-comté. references [1] t. w. clyne, s. c. gill. residual stresses in thermal spray coatings and their effect on interfacial adhesion: a review of recent work. journal of thermal spray technology 5(4):401–418, 1996. doi:10.1007/bf02645271. [2] g. montay, a. cherouat, j. lu. residual stress analysis in crankshaft using the hole drilling method. in materials science forum, pp. 537–542. trans tech publications ltd., 2006. doi:10.4028/0-87849-414-6.537. [3] p. withers, h. bhadeshia. residual stress. part 2 – nature and origins. materials science and technology 17(4):366–375, 2001. doi:10.1179/026708301101510087. [4] s. kuroda, t. fukushima, s. kitahara. significance of quenching stress in the cohesion and adhesion of thermally sprayed coatings. journal of thermal spray technology 1(4):325–332, 1992. doi:10.1007/bf02647160. [5] s. kuroda, t. clyne. the quenching stress in thermally sprayed coatings. thin solid films 200(1):49–66, 1991. doi:10.1016/0040-6090(91)90029-w. [6] s. kuroda, t. clyne. the origin and quantification of the quenching stress associated with splat cooling during spray deposition. in 2nd plasma technik symposium, plasma technik, vol. 3, pp. 273–284. 1991. [7] l. pawlowski. the science and engineering of thermal spray coatings. john wiley & sons, ltd, 2008. doi:10.1002/9780470754085. [8] s. kuroda. properties and characterization of thermal sprayed coatings-a review of recent research progress. in thermal spray meeting the challenges of the 21st century, proc. 15th int. thermal spray conf. may, 1998, nice, pp. 539–550. 1998. [9] p. withers, h. bhadeshia. residual stress. part 1 – measurement techniques. materials science and technology 17(4):355–365, 2001. doi:10.1179/026708301101509980. [10] handbook of measurement of residual stresses. choice reviews online 34(04):34–2182–34–2182, 1996. doi:10.5860/choice.34-2182. [11] j. mougin, t. l. bihan, g. lucazeau. high-pressure study of cr2o3 obtained by high-temperature oxidation by x-ray diffraction and raman spectroscopy. journal of physics and chemistry of solids 62(3):553–563, 2001. doi:10.1016/s0022-3697(00)00215-8. 46 https://doi.org/10.1007/bf02645271 https://doi.org/10.4028/0-87849-414-6.537 https://doi.org/10.1179/026708301101510087 https://doi.org/10.1007/bf02647160 https://doi.org/10.1016/0040-6090(91)90029-w https://doi.org/10.1002/9780470754085 https://doi.org/10.1179/026708301101509980 https://doi.org/10.5860/choice.34-2182 https://doi.org/10.1016/s0022-3697(00)00215-8 vol. 27/2020 residual stresses in cr2o3 coatings [12] j. mougin, n. rosman, g. lucazeau, a. galerie. in situ raman monitoring of chromium oxide scale growth for stress determination. journal of raman spectroscopy 32(9):739–744, 2001. doi:10.1002/jrs.734. [13] v. optasanu, e. bourillot, p. vitry, et al. high-resolution characterization of the diffusion of light chemical elements in metallic components by scanning microwave microscopy. nanoscale 6(24):14932–14938, 2014. doi:10.1039/c4nr04017a. [14] e. n. aybeke, s. ployon, m. brulé, et al. nanoscale mapping of the physical surface properties of human buccal cells and changes induced by saliva. langmuir 35(39):12647–12655, 2019. doi:10.1021/acs.langmuir.9b01979. [15] i. r. beattie, t. r. gilson. the single-crystal raman spectra of nearly opaque materials. iron(iii) oxide and chromium(iii) oxide. journal of the chemical society a: inorganic, physical, theoretical p. 980, 1970. doi:10.1039/j19700000980. [16] j. birnie, c. craggs, d. gardiner, p. graves. ex situ and in situ determination of stress distributions in chromium oxide films by raman microscopy. corrosion science 33(1):1–12, 1992. doi:10.1016/0010-938x(92)90014-t. [17] m. saka, y. ju, d. luo, h. abé. a method for sizing small fatigue cracks in stainless steel using microwaves. jsme international journal series a 45(4):573–578, 2002. doi:10.1299/jsmea.45.573. [18] r. zoughi. microwave non-destructive testing and evaluation. springer netherlands, 2000. doi:10.1007/978-94-015-1303-6. 47 https://doi.org/10.1002/jrs.734 https://doi.org/10.1039/c4nr04017a https://doi.org/10.1021/acs.langmuir.9b01979 https://doi.org/10.1039/j19700000980 https://doi.org/10.1016/0010-938x(92)90014-t https://doi.org/10.1299/jsmea.45.573 https://doi.org/10.1007/978-94-015-1303-6 acta polytechnica 27(0):42–47, 2020 1 introduction 1.1 residual stresses in thermal spray coatings 1.2 origins of residual stresses 1.3 purpose of the work 2 methods 2.1 materials and coating deposition 2.2 scanning microwave microscopy 2.3 raman micro-spectroscopy 2.4 x-ray diffraction 3 results and discussion 3.1 x-ray diffraction results 4 scanning microwave microscopy results 4.1 stress variation over centerlines 5 in-depth stress variation 6 raman spectroscopy results 6.1 stress variations over centerlines 7 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.29.0025 acta polytechnica ctu proceedings 29:24–28, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app bim – an application in geotechnical engineering ivan vaníček∗, jan pruška, daniel jirásko czech technical university in prague, faculty of civil engineering, department of geotechnics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: ivan.vanicek@cvut.cz abstract. bim model represents a significant step forward within the frame of the overall process of construction. this new approach goes from the assumption that up to date praxis, which is represented by different sets of 2d or 3d drawings of drawings, technical reports and calculations does not cover all problems of the above mentioned overall process of construction. basic principles of bim will be discussed, not only from the view of 3d models of the structure design. the utilization of the bim model during life time structure expectancy will be stressed from the initial phase (investigation, design) via the phase of structure execution, maintenance and ending with structure demolition at the end of life time expectancy. specificity of the geotechnical engineering is in this entire process fundamental, as each construction is in the interaction with ground. 3d model of the ground (geotechnical model) is therefore one of the basic individual parametric elements from which bim model consist. ground model is time dependant as geological profile and geotechnical properties are refinement during each phase of ground investigation as well during geotechnical structure construction. final 3d ground model together with 3d model of geotechnical structure represents a first significant step of the overall bim model. in the case of underground or earth structures such output can be primordial element of the bim model with parametric elements around it. finally some other possibilities or practical applications are mentioned. keywords: bim, geotechnical engineering, ground model, model of the realized geotechnical structure, sustainable construction. 1. introduction bim – building information modelling (management) as a tool for the process of optimizing construction activities is connected over the most recent period with a more general digitalization evident throughout all industries. the term průmysl 4.0 (industry 4.0) or stavebnictví 4.0 (building 4.0) is now standard for most elements of building engineering. as from now more attention is being devoted also in civil engineering and some significant czech investors and construction companies are demanding that they receive data in the form of the bim model in order to be able to digitalize the process of construction. however the main condition for the application of the bim model in the wider profession is a clear understanding of bim principles. there are many different aspects to a definitive bim. the first one, that bim is simply a 3d model of the engineering structure, is not correct as it does not take into account up to date instruments of information modelling concerning engineering structures and modern communication methods. the most widely used definition was specified by the american national committee for the bim standard: “building information modelling (bim) is a digital presentation of physical and functional characteristics of a facility. a bim is a shared knowledge resource for information about a facility forming a reliable basis for decisions during its life-cycle, so defined as existing from earliest conception to demolition” [1]. in other words bim can be understood as a process for optimization of the engineering structure right through conception, design, execution and operation. by this approach the 3d computer model is a fundamental part of bim but however is enlarged by additional information on such matters as time, financial expenses, structure maintenance etc – with new so called 4d, 5d, and generally nd models. 2. the situation in the czech republic the decree of the government no. 82: “a concept for the implementation of the bim model in the czech republic” instigated september 25 2017 can be taken as the basis of bim acceptance. the main driver is the ministry of industry and trade which accredited the czech agency for standardization (cas) responsible for the realization of this governmental decree. some results of cas activities are described in [2]. from this analysis some basic reflections, facts and expected developments can be mentioned. from the reflections there can be cited “bim is about change – respectively processes, ideas, working habits, and “ the improvement of cooperation, information sharing along with an interrelationship cultivation that helps the higher effectiveness of engineering structures projects”. from the actual facts, six 25 http://dx.doi.org/10.14311/app.2020.29.0025 http://ojs.cvut.cz/ojs/index.php/app i. vaníček, j. pruška, d. jirásko acta polytechnica ctu proceedings working groups can be mentioned and some results are presented at www.koncepcebim.cz. activities of cas are coordinated in a european framework by positive participation in cen tc 442 – building information modelling (bim). the advantages of the bim are now primarily emphasized with respect to the practical application for engineering facilities of buildings, e.g. in articles [3] and [4]. the authors of the second paper dealing with building of a hospital in náchod stress the ratio of the construction to the operating charges for the expected life time expectancy of the building which they estimate to be 20 : 80. at the same time of course they are stressing the main idea: “the more quality we buy at the beginning, the less we spend later on”. naturally, this should be valid also for civil engineering structures such as transport infrastructure. when the transport engineering structure is made to be more robust, then less should be needed in the future for maintenance, reconstruction or on the restoration of services after natural hazards such as e.g. floods, rock falls; as any restraint of services can be included into additional expenses. the third note refers to the activities of sfdi – governmental fund of transport infrastructure which is responsible for bim implementation for transport engineering. procedures of sfdi are very important with the respect to the above-mentioned governmental decree according to which all significant funded projects disbursed from the government budget should be contracted within the bim regime from 2022. sdfi, currently in september 2019, approved the provisional version of the document “directive for bim pro structures of transport infrastructure – data standard – for pdps – project documentation of structure realization (execution)”. in this provisional version attention is devoted to the project documentation for structure execution, and therefore up to some extent it is outwith the main aim of the bim – to interconnect previous steps of the project documentation, even when some parts of the dur (documentation for planning inquiry), dsp (documentation for building permit) are mentioned in annexes 1 and 2 dedicated to road and railway structures. a “directive” can specify rules for the collection of data for bim in such a way that it can be used by owners, designers, contractors in all phases of design, execution and transport infrastructures maintenance. the data standard is based on the open data format ifc, see čsn en iso 16739 (730100) from april 2017. the ifc data format is playing the role of the coordination model of construction, which is composed of the individual construction sets, operational sets or engineering objects. this co-ordination model is used for the mutual coordination of the fragmental models, for the detection of collisions, in all construction display, for display of individual phases of the construction process and also for creation of cross sections. the co-ordination model is in principle an individual set, which contains the fragmental models. 3. bim and geotechnical engineering geotechnical engineering has an indispensable role in bim , as all structures are in some interaction with ground and this interaction is provided by foundation structures – fig. 1. the term structure is composed not only of the upper structure but also of foundation structures, including ground which is influenced by the total construction and includes a view of stresses, deformation or permeability. fragmental models for which the geotechnical engineering is responsible should therefore consist of the ground model, the geotechnical design model, the calculation model and also by a model of the real execution of the geotechnical structure with its connection to the ground model. these fragmental models from which all others models proceed can bring a significant recognition of the role of geotechnical engineering in the whole construction process. at the same time these fragmental models can lead to a more significant utilization of the basic principles of, and demands on, geotechnical structures, which are in europe specified in sets of eurocodes, initially in eurocode (ec 7) geotechnical design. figure 1. interaction of structure and subsoil all the above mentioned is valid for foundation structures, see fig. 2.a., but similarly, if not more significantly, for other geotechnical structures such as earth structures (for transport, water or environmental engineering) or underground structures (tunnels, underground repositories, metro. . . ) as these geotechnical structures can even be the main element of the solved construction, fig. 2.b. the significance of fragmental geotechnical models stands front rank as the main project idea in the phase of alternative solutions of the construction situation (placements) that should arise from the first phases of the ground (geological) model. this first phase of the geotechnical investigation is denoted as desk study and uses all previous information, firstly different geological maps, then engineering geological maps, hydrogeological maps , more generally from the 26 vol. 29/2020 bim – an application in geotechnical engineering figure 2. the position of geotechnical bim model for various geotechnical structures. ad a) geotechnical structure creates a basic bim model. ad b) gs creates an individual parametric elementaround basic bim model. set of so called geo-environmental maps, as this set includes also information about raw materials, about quality of underground water which should be protected etc. their utilization is also for the eia process – environmental impact assessment – for the phase of the evaluation of different variants of the transport infrastructure lines. the ground model is step by step improved so as to be able to fulfil demands on the individual stages of the design and will be terminated by the model which was verified during the phase of structure execution. the final version can be used for whole life time expectancy should any new interaction start. 3.1. foundation structures the 3d model of the ground together with the 3d model of the foundation structure can be used for verification of the basic rules which are specified in the eurocode 7 geotechnical design and involve: • to what degree the investigation methods respect the significance of the geotechnical structure and the risk with which any failure is connected – which is in the ec 7 expressed by 3 different geotechnical categories; • whether the depths of the investigation works are catching the maximum depths influenced by the proposed construction; • whether the range of the proposed and completed field and laboratory tests make possible their credible evaluation; • what differences there are between the measured values of geotechnical parameters, and which were finally used during the design as characteristic values, and with the values obtained during structure execution. similarly this is valid for step by step improvement of boundaries between individual lithological layers. as the bim model is time dependant there is the possibility to intervene at any time to correct potential shortcomings. similarly this is valid for geotechnical structures for which limit states are very sensitive to time, as from the uls with respect to short-term bearing capacity (stability) so on for long-term bearing capacity (stability) and then so for the sls verification. therefore, geotechnical structures get benefit from the monitoring plan not only during the phase of construction but also for the phase of construction of the upper structure, as well as during the period of warranty, to confirm that all demands are satisfied. the monitoring plan or plan of maintenance can create an independent fragmental model. the bim model is also able to control the schedule of the construction process, e.g. piles with additional information about pile execution, testing and a quality of material confirmation. 3.2. earth structures as the set-up of the bim model is very significant for transport infrastructure our attention is focused on these earth structures in transport engineering. timing specification is shown in fig. 3. for individual parts of the geotechnical design report (gdr). there are two geological models, the first one for subsoil and the second one for borrow pit, and from which material will be used for fill. similarly tests will be carried out for these two basic materials. for embankment, the quality of subsoil is not controlled during execution, and attention is devoted to the control of deposited and compacted material from the borrow pit. in principle there should be two levels of control, for compaction and for material properties (strength, deformation, filtration parameters) as these were used during the design. however, the possibilities in the second direction are very limited. the 3d model of subsoil together with the constructed earth structure (embankment) is shown in fig. 4. there are also supplementary cross-references on the description of investigation points (boreholes) on measure properties, and on the final selection of characteristic values of geotechnical parameters. similarly, as for foundation structures, the bim model can be used for approval on whether the model of subsoil is sufficient specifically with respect to the geotechnical category into which the designed geotechnical structure is falling. 3.3. underground structures (tunnels) as the bim model is proposed for total life time expectancy a significant role is present also for the phase of structure preparation (study). this is so especially with the respect to traffic line selection, the selection of technology for tunnel performance, a statement about material demands, and financial cost estimation. for embankment or cuts the estimation is relatively easy compared to the estimation for tunnels. the main reason is that during a tunnel execution unexpected geological and geotechnical conditions can be discovered because during the actual geotechnical investigation only a limited volume of the rock massif can be in fact be subject to this investigation. in this 27 i. vaníček, j. pruška, d. jirásko acta polytechnica ctu proceedings figure 3. geotechnical design report figure 4. vision of 3d model of earth structure within bim case the proposed bim model is only approximated and any final modification according to the real geological and geotechnical conditions should be the result of the last phase of performance. however, this last bim version is very important for the phase of tunnel operation and maintenance – for facility management (fm). as in bim , so also fm is a multi-criteria process. sometimes fm is denoted in bim as a sixth nongeometrical dimension – 6d bim. the ita – ites (international tunnel association) devotes great attention to bim at present. this is namely with respect to tunnel information modelling (tim) with the main aim to approve and support closer cooperation in the context of information management [6]. 4. summary bim is a long-term process which requires the preparedness of the people involved in, and familiar with, 28 vol. 29/2020 bim – an application in geotechnical engineering the new methods and processes which bim is offering. with wider application of bim in building and civil engineering new and more user-friendly software is appearing in this market. it is obvious that in the future transport infrastructures projects will almost certainly be realized involving bim and information management. the rules for multi-disciplinary cooperation of individual subjects are already partly included in the standard of ifc. however, these standards are not applicable for some civil engineering structures such as tunnels. that is the reason why under the umbrella of ita-ites a new working group was established, namely for above mentioned tim. nevertheless, in the field of geotechnical engineering bim can bring many new positive features, predominantly from the viewpoint of limitation of the risk with which geotechnical structures are connected. creation of the 3d model of subsoil together with the 3d model of geotechnical structures (either earth, foundation of underground) will help bring a better recognition of the geotechnical profession, especially its role in the whole process of construction. a great possibility exists with connection to the basic code for geotechnical structures – eurocode 7 geotechnical design. the creation of individual models is the common endeavour of ec 7. as for the standpoint of interconnection of individual phases and also for their control, that basic recommendation of the ec 7 appears satisfied. a full exploration of bim is duly justified given the basic aim defined for the bim which is that it should be a reliable basis for the decisionmaking process during a full life time structure expectancy – from the early phase of conception up to the structure demolition. after that it would be easier to implement simply from principles of sustainable construction. however, it is very important to define the principles of sustainable construction from the earliest phase. for example, the effort directed towards lower land consumption (greenfield saving) should be defined at the beginning as it is connected also with transport line selection. also, the possibility of land savings due to application of steeper slopes which are reinforced should start as soon as possible. however, the principle of sustainability is connected also with savings of energy and natural materials as for example in the case of reinforced retaining structures – during the phase of final decision about their type – as this principle can play significant role, additionally with respect to the c02 path. similarly it is valid for application of large volume waste materials into embankment such as flying ash, construction and demolition waste etc. instead of using natural soil. the transfer of the currently used 2d subsoil model into 3d is not such a great model and the first signs of this are becoming recognisable. similarly this is valid for design verification, with continuous transfer from 2d models into 3d models, even if the transport infrastructure where one dimensionality prevails, is more typical for 2d. therefore, more positively discussed is the transfer of “our” now typical geotechnical models into the coordination model of bim for which the main coordinator is responsible. this transfer demands a good orientation to code čsn en iso 16739 industry foundation classes (ifc) for data sharing in the construction and facility management industry (iso 16739:2013), especially, with respect to the fact that it is an open neutral standard data model, the transfer should not be a great problem – certainly after acquaintance develops confidence. the main coordinator will guarantee that this data model will be free of charge. a final note merits focus on the utilization of bim during life time expectancy. a basic model of subsoil together with a 3d model of geotechnical structure will be utilizable in the case of any new interaction. interaction with a new construction affecting the old one or due to any natural hazards or indeed any manmade hazards such as different types of accidents, e.g. a lorry with chemical substance, can allow a quick reaction to protect subsoil contamination. acknowledgements the paper was prepared with financial support from the programme of the centre of competence of the technology agency of the czech republic (ta čr), no. of project te0120168. references [1] us national bim standard project committee. [online]. [28. 8. 2019]., available at: https://www.nationalbimstandard.org/faqs#faq1. [2] j. nechyba (2018). vládní koncepce bim – rok po jejím schválení. zprávy a informace čkait, 5/2018. [3] p. kuba, d. wallensfels (2019). nová vodní linka účov: cesta od projektování 2d k projektování 3d a bim. stavitelství 04/19. [4] j. slánský, p. bečička. (2019). využití bim na příkladu projektu nemocnice náchod. stavitelství 09/19. [5] m. žibert. bim in tunnelling – implementation of bim methodology to trans alpine tunnel karavanke. brisbim gathering, august 2017, [online]. [28. 8. 2019]. available at: http://brisbim.com/wp-content/uploads/2017/08/ 170822_ic_bim-brisbim-kartu_mzr.pdf. [6] working group 22: information modelling in tunnelling. international tunnelling and underground space association [online]. [28. 8. 2019]. available at: https://about.ita-aites.org/wg-committees/ working-groups/276-ita-active-working-groups/ working-group-22-new-2017-informationmodelling-in-tunnelling. 29 https://www.nationalbimstandard.org/faqs#faq1 http://brisbim.com/wp-content/uploads/2017/08/170822_ic_bim-brisbim-kartu_mzr.pdf http://brisbim.com/wp-content/uploads/2017/08/170822_ic_bim-brisbim-kartu_mzr.pdf https://about.ita-aites.org/wg-committees/working-groups/276-ita-active-working-groups/working-group-22-new-2017-information-modelling-in-tunnelling https://about.ita-aites.org/wg-committees/working-groups/276-ita-active-working-groups/working-group-22-new-2017-information-modelling-in-tunnelling https://about.ita-aites.org/wg-committees/working-groups/276-ita-active-working-groups/working-group-22-new-2017-information-modelling-in-tunnelling https://about.ita-aites.org/wg-committees/working-groups/276-ita-active-working-groups/working-group-22-new-2017-information-modelling-in-tunnelling acta polytechnica ctu proceedings 29:24–28, 2020 1 introduction 2 the situation in the czech republic 3 bim and geotechnical engineering 3.1 foundation structures 3.2 earth structures 3.3 underground structures (tunnels) 4 summary acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0062 acta polytechnica 27(0):62–66, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app hardness of hybrid pvd-pecvd w-c:h coatings vs. substrate type františek lofaj∗, margita kabátová slovak academy of sciences, institute of materials research, watsonova 47, 040 01 košice, slovakia ∗ corresponding author: flofaj@saske.sk abstract. the effects of substrate on the measurement of hardness (and indentation modulus) penetration depth profiles were investigated on hybrid pvd-pecvd w-c:h coatings made with the additions of c2h2 or ch4 using high power impulse magnetron sputtering (hipims) method. the substrates involved 100cr6 bearing steel, al2024 aluminum alloy and (111) si wafer. although no clear influence of the substrate type on hardness and indentation modulus of the coatings at different acetylene and hydrogen additions was observed, when both measured properties were combined in hiteit ratio, significant shift of hit/eit ratio dependencies obtained on al alloy vs. steel and si wafer substrates occurred. thus, the effect of substrate type on the measurementlevaluation was confirmed. keywords: hardness, indentation modulus, substrate effect, w-c:h coating. 1. introduction tribological properties of hard coatings, especially their wear resistance, coefficient of friction (cof) depend on large number of parameters including loading and environmental conditions but also material system used for the coating, its structure, surface roughness and such properties as hardness, elastic modulus, and others. it was reported that the wear properties of hard coatings more strongly correlate with h/e ratio than just with the hardness itself [1]. moreover, it was also found out that the h/e and/or h3/e2 ratios are related to the level of plasticity and can be used to estimate changes in the toughness of the coatings [2–4]. this is of significant interest because toughness is among the main parameters, besides hardness, young’s modulus, cof, wear, scratch, oxidation and corrosion resistance of the coatings for engineering applications. the measurement of mechanical properties, involving hardness, elastic modulus and toughness in the coatings with the thickness in the range of few micrometers is a challenge even by nanoindentation because of the inevitable influence of the substrate on the measurement of the coating properties. to reduce this influence, the indentations depths used for the determination of coating hardness should be substantially smaller than the thickness of the coating. physical explanation is that the stress field under the indenter, corresponding to the zone with the stresses above the yield stress (zone of plasticity), has to be fully confined in the coating. so called “10 % (of the coating thickness) rule” is often applied to determine such maximum depth and true hardness of the coating, respectively [5–10]. another approach to eliminate substrate influence and to determine true coatings properties is to use hardness (and elastic modulus) – indentation depth profiles obtained from continuous multi-cycle (cmc) or continuous stiffness measurements (csm) [11]. the coating properties correspond to the maximum or plateau on the corresponding depth profile at the depths smaller than 10 % of the coating thickness. however, the liability of 10 % rule is limited by the accuracy of the measurement at coating thicknesses below 1 µm related to tip surface area calibration function, by indenter tip radius [12], surface roughness, pile-up or sink-in behavior [13], residual stresses as well as by the ratio between the yield stresses and the elastic moduli of the coating and the substrate, respectively [9, 14]. the measurements of indentation modulus are much more sensitive to substrate properties than hardness because the elastic stress field under the indenter would be much larger than the zone of plasticity and therefore, 10 % rule may not be sufficient. thus, the problem of proper measurement of hardness simultaneously with the elastic modulus in thin hard coatings persists and it is even emphasized by the applicability of the h/e ratio for toughness and wear resistance estimations. in our earlier works, the influence of indenter tip radius on the hardness – depth profiles has been analyzed using a combination of the experimental and finite element modelling approaches [12]. however, the influence of the young’s modulus of the substrate on the nanoindentation measurements was not investigated. therefore, the aim of the present work is to investigate the influence of the substrate properties on the measurements of hardness, and indentation modulus in thin hard coatings used for engineering applications. the measurements were performed on hipims (high power impulse magnetron sputtering) w-c and w-c:h coatings deposited using hybrid pvd-pecd (physical vapor deposition – plasma enhanced chemical vapor deposition) process with different acetylene and hydrogen additions on three different substrates. 62 http://dx.doi.org/10.14311/app.2020.27.0062 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 coatings vs. substrate type hardness figure 1. load – indentation depth curves from the nanoindentation tests on the hipims w-c:h coatings deposited with 2 sccm c2h2 on si wafer, 100cr6 bearing steel and al 2024 alloy substrates – a) the depth profiles of hardness – b) and indentation modulus – c). 2. experimental material and methods the studied w-c:h coatings were simultaneously deposited on three different substrates: polished tempered bearing steel 100cr6 discs (diameter – 25mm, thickness – 3mm ), polished al2024 aluminum alloy discs with similar size and on approximately 1cm2 polished (111) si wafer fragments with ∼ 200nm cr bond layer. the deposition was carried out at a minimum working pressure of 0.5 pa in a magnetron deposition system (cryofox discovery 500, polyteknik, denmark) using high power impulse magnetron sputtering (hipims) source at 350 w average power achieved at 2.62% duty cycle related to the frequency of 150hz and impulse length of 175 µs. a set of 14 coatings with variable additions of c2h2 precursor (0, 2, 4, 6, 8 sccm) and (0, 5, 10, 20 sccm) into ar atmosphere was prepared for the study. the thicknesses of the coatings were in the range from 0.7 µm up to almost 3 µm depending on the amount of precursor gas additions [15–18]. the nanoindentation tests were carried out on a nanoindenter (model g200, agilent, usa) using a diamond berkovich tip in csm mode with the preset maximum penetration depth of 1000 nm, strain rate of 0.05 s−1, frequency of 45 hz and amplitude of 2 nm on a set of 16 indents. the hardness and indentation modulus depth profiles were obtained by averaging the results of at least 10 measurements. the young’s modulus and poisson’s ratio of the diamond tip used for calculations were etip = 1141 gpa and νtip = 0.07 , respectively. the indenter tip was calibrated on a fused silica reference sample. the poisson’s ratios of the coatings of v = 0.25 was used for the evaluation of the corresponding properties. 3. results and discussion 3.1. substrate effects on hardness and indentation modulus in hipims w-c:h coatings fig. 1a 1c illustrate the differences between load indentation depth curves, hardness – indentation depth and indentation modulus – depth profiles in the studied w-c:h coating deposited with the addition of 2 sccm c2h2, respectively. the load – displacement curve in the coating on al alloy differs significantly from those on si and steel substrates, which were practically identical, at the displacements above 400 nm 63 františek lofaj, margita kabátová acta polytechnica figure 2. the comparison of the hardness, hit − a) and indentation modulus, eit − b) in hipims w-c:h coatings deposited at different acetylene (and hydrogen) additions on three types of substrates. the data-points corresponding to 5, 10 and 20 sccm h2 additions were intentionally shifted by 0.2, 0.4 and 0.4 sccm respectively, for better visibility. (fig. 1a). however, this difference is not important for the determination of the hardness and indentation modulus, because they have to be determined from the indentation depths within the range of 100 300 nm due to the earlier mentioned 10 % rule. fig. 1b shows the depth profile of hardness: on si and steel substrates, very wide plateau at their maxima of around 20 gpa extending up to almost 500 nm depth can be seen. in contrary, the influence of substrate is much stronger in the case of measurement on al substrate: the onset of hardness decrease occurred already at 250 nm. such difference is a natural consequence of lower young’s modulus of al alloy in comparison with steel and si wafer. however, the maximum of the hardness profile obtained on al alloy overlapped with those on the other two substrates and essentially no difference of the substrate on hardness can be seen. however, the situation was different in the case of indentation modulus profiles. a plateau at around 225 gpa extending up to 500 nm can be seen on steel substrates but the measurement on si wafer resulted in a maximum plateau at around 220 gpa only up to ∼ 300 nm. the effect of al alloy substrate was much more pronounced: the maximum was around 175 gpa and it extended only to 150 nm indentation depth. the values of hit and eit determined from the corresponding depth profiles according to the above mentioned way are summarized in fig. 2a and 2b, respectively. they reveal number of effects and differences. they involve the effects of substrate type but also of acetylene and hydrogen additions during hybrid pvd-pecvd. the pvd process without additional precursor gases resulted in (possibly over-stoichiometric [18]) wc coatings exhibiting much lower values of hardness (22 gpa) and modulus (< 220 gpa) in the coatings deposited on al alloy substrates than on si and steel substrates (26 – 29 gpa and 280 – 340 gpa, respectively). it is known from earlier works on dc magnetron sputtered and hipims w-c:h coatings [17, 18] that the structure of the pvd coatings without additions of acetylene may vary from nano-columnar to nanocrystalline, which resulted in the variations of hardness from 34 gpa to 28 gpa, respectively. the additions of acetylene caused gradual increase of the content of hydrogenated free carbon phase and a transition from nanocomposite to amorphous structure accompanied by a decrease of hardness [16–18]. the substrate itself may also play significant role in the formation of the coating structure and its morphology [15]. without detail structure observations, it is difficult to decide, which of the mentioned effects could be responsible for the changes of hardness and indentation modulus in the studied w-c coatings. as expected, the additions of acetylene resulted in gradual decrease of these properties. there were certain variations among the values obtained on various substrates, but they were within the scatter bars of the measurements and the differences decrease, especially in the case of elastic moduli (fig. 2b). apparently, the amorphization of the structure at high acetylene additions seems to result in the unification of the properties and has stronger effect than the possible influence of the substrate. thus, the differences in coating properties deposited on different substrates seem to be statistically significant only in the cases without or with very small acetylene additions. however, they may interfere with the effects of coating structure and therefore, cannot be fully attributed to the substrate effect. figs. 2 also show the influence of hydrogen additions; it should be noted that the data-points corresponding to 5, 10 and 20 sccm h2 additions were intentionally shifted by 0.2, 0.4 and 0.4 sccm from the position of the corresponding acetylene addition to show the tendency of the changes without overlap. it can be seen that the hydrogen additions have different effects on hit and eit in pvd w-c coatings and in hybrid pvd-pecvd w-c:h coatings made with acetylene additions. in the first case, hydrogen additions resulted in the increase of hardness and elastic modulus 64 vol. 27/2020 coatings vs. substrate type hardness whereas they were decreased in the coatings made with 4 sccm and 6 sccm acetylene additions regardless of the substrate. moreover, the corresponding values principally followed the dependencies described for the coatings without hydrogen additions and agree with the above conclusions concerning exclusivity of w-c coatings compared to w-c:h coatings. the above results indicate only a small influence of substrate type on hardness and indentation modulus. however, they are measured simultaneously and it is reasonable to check the simultaneous effect of both properties on different substrates. thus, the results from fig. 2a and 2b were replotted as a function of hit/eit ratio on acetylene (and hydrogen) additions in fig. 3. the obtained plots show that the acetylene and hydrogen additions usually result in gradual hit/eit increase and that the difference between substrates was strongly magnified. although the differences between hit/eit in the coatings on 100cr6 steel and si substrates were negligible in the studied additions range, the corresponding values on al alloy substrates followed the same tendency but they were by 0.02 – 0.025 higher. since both hardness and indentation modulus decrease with the acetylene addition increase, the increase of hit/eit has to result from faster decrease of indentation moduli. another consequence of the shift of the dependencies between steel, si and al alloy substrates was that all hit/eit values obtained on al alloy were above the plasticity limit of 0.1 while the identical coatings deposited on si and steel substrates at 0 and 2 sccm acetylene (and hydrogen) additions would be below that limit. obviously, such systematic differences in hit/eit on the same coatings do not have real physical reason. they have to be related to the influence of the substrate type on the measurement method and/or data evaluation rather than on true hardness and indentation modulus changes. it is a clear indication that the substrate type definitely affects the nanoindentation measurements. the question is what is the main parameter controlling the influence of the substrate. young’s moduli of steel and si substrates are rather similar (210 gpa vs. 193 gpa) while that of al alloy is considerably lower (∼ 73 gpa) and this difference is reflected in the difference obtained fig. 3. however, the yield strength and hardness of substrate materials should play a role because they control plasticity of the substrate 4. conclusions the investigations of the influence of the substrate type on the measurements of hardness and indentation modulus in hipims w-c and w-c:h coatings with different acetylene and hydrogen additions on si wafer, 100cr6 steel and al 2024 aluminum alloy substrates revealed: • significant influence of substrate type on hit and eit in pvd w-c coatings deposited without the 0 2 4 6 8 0.075 0.100 0.125 0.150 hipims w-c:hcoatings 5,10,20 h 2 5,10,20 h 2 h it /e it ra tio xc 2 h 2 (+ yh 2 ) flow, sccm alsubstrate sisubstrate fe substrate 5,10,20 h2 al 2024 substrate sisubstrate 100cr6 substrate figure 3. the dependence of hit /eit ratio in hipims w-c:h coatings on the substrate type at different acetylene (and hydrogen) additions. note that the data-points corresponding to 5, 10 and 20 sccm h2 additions to the corresponding acetylene flows were intentionally shifted to emphasize the role of hydrogen additions. additions of acetylene. however, these differences may overlap with the influence of substrates on the coating structure. • an enhanced scatter of the measured properties in hybrid pvd-pecvd w-c:h coatings with the additions of acetylene on the studied substrates. the differences between the substrates were within the scatter of measurement. • when the effects of substrate type on hardness and indentation modulus were combined in hit/eit ratio, significant shift between the dependencies obtained on al alloy vs. steel and si wafer substrates was obtained. thus, the effect of substrate type on the measurement/evaluation was confirmed. • the effect of substrate properties on nanoindentation measurements cannot be related only to the difference in young’s moduli of the substrates and further investigations are required to find the way to eliminate this influence. the above conclusions imply that the reliable evaluation of the wear resistance and fracture toughness of thin coatings based on hit/eit ratio requires the elimination of the influence of the substrate type on hit/eit ratio. acknowledgements the support provided by the projects apvv-17-0320, apvv 17-0059, apvv-17-0049 and vega 2/0017/19 is acknowledged. the equipment used in the work was acquired from the projects “research centre of advanced materials and technologies for recent and future applications” promatech, 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https://doi.org/10.1590/0366-6913201965s12577 https://doi.org/10.1016/j.surfcoat.2019.07.078 https://doi.org/10.1016/j.ijrmhm.2020.105211 acta polytechnica 27(0):62–66, 2020 1 introduction 2 experimental material and methods 3 results and discussion 3.1 substrate effects on hardness and indentation modulus in hipims w-c:h coatings 4 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0610 acta polytechnica ctu proceedings 33:610–616, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague effect of using biomass fly ash on the concrete sustainability elisabete rodrigues teixeiraa, ∗, aires camõesb, fernando g. brancoc a university of minho, isise, department of civil engineering, guimarães, portugal b ctac, department of civil engineering, school of engineering, university of minho, campus de azur’em, 4800-058 guimarães, portugal c university of coimbra, isise, department of civil engineering, coimbra, portugal ∗ corresponding author: b8416@civil.uminho.pt abstract. biomass fly ash has been studied has a partial cement material, since this material has a positive effect on concrete properties. even though, some mixes with bfa presents a positive benefit, they need to be economically competitive and present a good environmental performance. so, the analyse and comparison of concrete that uses bfa as raw material substitution in terms of environmental impacts related to the production of conventional concrete. one of the best approaches to develop this type of study is to use the life cycle assessment (lca) method. this method quantifies both the input flows, such as energy, water and materials, as well as the output flows, such as co2 emission, solid wastes and liquid wastes. based both on the abovementioned context and methodological approach, a quantification and comparison of potential environmental impacts resulting from the production of 1 m3 of concrete was made, using different types of binder and quantities of cement substitution. keywords: life cycle analysis, sustainability, wastes. 1. introduction biomass fly ash (bfa) has an effect on concrete properties, presenting similar results to concrete with traditional pozzolanic materials such as coal fly ash [1]. moreover, the analysis of the potential environmental impacts related to the production of a concrete that uses bfa as raw material substitution is mandatory. this analysis needs to be compared with the production of plain cement concrete. one of the best approaches to develop this type of study is using the life cycle assessment (lca) method [2]. this method enables the quantification of the potential environmental impacts of products or services. it quantifies both the input flows, such as energy, water and materials, as well as the output flows, such as co2 emission, solid wastes and liquid wastes [3]. lca allows estimating the potential impact on humans and nature and enables identifying areas with improvement potential [3]. based both on the abovementioned context and methodological approach, a quantification and comparison of potential environmental impacts resulting from the production of 1 m3 of concrete was made, using different types of binder: i) portland cement; ii) portland cement and cfa or/and bfa, iii) portland cement and cfa and hl; and iv) portland cement, cfa, bfa and hl. this work was presented for the defence of the first author phd degree and it is presented in [4]. 2. methods and methodology 2.1. goal and scope the main goal of this study was to evaluate the environmental performances of various concrete formulations using biomass fly ash as a cement replacement material (table 1) [4]. the method used in this study followed the phases of an lca. the use of lca is done according to the iso 14040 standard, which provides a consensual framework, terminology and methodological phases [3]. the implementation of this method is based on four major phases: i) goal and scope definition; ii) inventory analysis; iii) impact assessment; and iv) interpretation [3]. the goal and scope express the purpose, objectives, product system, boundaries and functional unit. in the inventory analysis, the data necessary to analyse the life cycle of the product is collected. in the impact assessment, the life cycle inventory (lci) flows are classified, characterized and normalized, using one of many possible life cycle impact assessment (lcia) methodologies to estimate the potential environmental impacts. the last phase, interpretation, is very important to : i) identify, quantify and evaluate the information that results from the last two phases; ii) communicate the information in a correct way; and iii) recommend improvements within the analysed system [3, 4]. the comparative analysis and the aggregation of indicators were developed using the multicriteria decision support methodology for the relative sustainability assessment of building technologies (mars-sc) [5]. the mars-sc methodology is 610 https://doi.org/10.14311/app.2022.33.0610 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 biomass fly ash on the concrete sustainability concrete mix %wt w/b cement bfa cfa hl ref 100.0 0.0 0.0 0.0 0.50 cfa20 80.0 0.0 20.0 0.0 0.50 bfa20 80.0 20.0 0.0 0.0 0.50 cfa40 60.0 0.0 40.0 0.0 0.50 bfa40 60.0 40.0 0.0 0.0 0.50 cfa50 50.0 0.0 50.0 0.0 0.50 cfa50b 50.0 0.0 50.0 0.0 0.35 bfa50 50.0 50.0 0.0 0.0 0.5 cfa60 40.0 0.0 60.0 0.0 0.5 bfa60 40.0 60.0 0.0 0.0 0.5 cfa49.5hl0.5 50.0 0.0 49.5 0.5 0.35 cfa49.5bfa0.5 50.0 0.5 49.5 0.0 0.35 cfa48.8bfa1.3 50.0 1.3 48.8 0.0 0.35 cfa45bfa5 50.0 5.0 45.0 0.0 0.35 cfa48.8bfa0.6hl0.6 50.0 0.6 48.8 0.6 0.35 table 1. binder fraction and water/binder ratio used in the concrete formulations [4]. based on three groups of sustainability categories: environmental, functional and economic [6]. since this research aimed at assessing the environmental performance of different concrete formulations, only the environmental category of mars-sc was considered. the mars-sc methodology is processed in five steps: i) definition of the sustainability indicators; ii) quantification of the indicators (including the life cycle inventory); iii) normalization of the indicators; iv) aggregation of the indicators; and v) sustainable score calculation and global assessment [5]. 2.2. system boundaries and functional unit the boundaries of this work mark the embodied environmental impacts (cradle-to-gate) of the different concrete compositions as well as the environmental impacts that result from the transportation of the materials to the concrete plant and their mixing. the choice of limiting the study to the cradle-to-gate stage is justified by the fact that, in the studied compositions, the use and disposal of concrete will result in similar environmental impacts [4]. the declared functional unit is dependent on the goal of life cycle analysis and therefore constitutes 1 m3 of concrete. 2.3. inventory analysis to quantify the sustainability indicators, it is necessary to first develop the inventory analysis [6]. the inventory is used to quantify the inputs (e.g. energy, materials and chemical) and outputs (e.g. emissions and wastes) of the product system [7].) table 2 shows the inventory of the materials and transportation considered for each concrete formulation [4]. this inventory took into consideration the specific context of the portuguese concrete industry. the life cycle analysis software simapro 7.3.3 was used to facilitate the quantification of the impact categories. in this study, the specific consumption of raw materials, energy and fuels and the emissions released to air, water and soil during the cement production of an important portuguese cement plant, located in the south of portugal, was considered. the data used are described in the public environmental declaration [8] of this cement plant. for this research, it was considered that this plant supplied the cement used for the preparation of the different concrete formulations. it was necessary to quantify the impact categories, since the environmental declaration did not cover all impact categories necessary for this study, being limited to those mandatorily declared according to portuguese environmental legislation [4]. using the inventory listed in the environmental declaration, the simapro software was used to assess the potential environmental impacts of the used portland cement. regarding each type of fly ashes, it was necessary to make the allocation of flows of the power plant according to the place of production. allocation is necessary in the case of joint coproduction, where the processes cannot be sub-divided, as is the case in fly ashes production [9]. allocation shall respect the main purpose of the processes studied, appropriately allocating all relevant products and functions. since the main purpose of a thermal power plant is to produce electricity and since the difference in revenue between the electricity and the fly ashes is high, it is not possible to use an allocation process based on physical proprieties (e.g. mass and volume). therefore, the allocation process used in this research was based on economic values [4]. due to the environmental report [10] from a major portuguese coal power plant (located in the centre 611 e. r. teixeira, a. camões, f. g. branco acta polytechnica ctu proceedings formulations pc* gravel sand water sp* cfa bfa hl material input (kg) ref 350.0 969.2 738.1 200.0 0.0 0.0 0.0 0.0 cfa20 280.0 982.1 756.7 149.5 0.0 70.0 0.0 0.0 bfa20 280.0 985.3 759.1 149.4 0.0 0.0 70.0 0.0 cfa40 210.0 973.0 749.6 149.7 0.0 140.0 0.0 0.0 bfa40 210.0 979.3 754.5 149.5 0.0 0.0 140.0 0.0 cfa50 175.0 968.4 745.1 149.8 0.0 175.0 0.0 0.0 cfa50b 175.0 1044.3 804.5 88.4 8.8 175.0 0.0 0.0 bfa50 175.0 976.3 752.2 149.6 0.0 0.0 175.0 0.0 cfa60 140.0 963.8 742.5 149.9 0.0 210.0 0.0 0.0 bfa60 140.0 973.3 749.9 149.7 0.0 0.0 210.0 0.0 cfa49.5hl0.5 175.0 1049.0 768.6 91.5 5.3 173.5 0.0 1.9 cfa49.5bfa0.5 175.0 1049.0 795.2 91.5 11.4 173.5 1.9 0.0 cfa48.8bfa1.3 175.0 1050.1 842.8 35.5 11.4 170.6 4.4 0.0 cfa45bfa5 175.0 1049.5 838.9 42.4 3.9 160.0 17.8 0.0 cfa48.8bfa0.6hl0.6 175.0 1067.2 832.9 36.9 2.5 170.6 2.2 2.2 transportation (tkm) ref 14.5 357.6 272.4 − − − − − cfa20 11.6 362.4 279.2 − − 12.4 − − bfa20 11.6 363.6 280.1 − − − 10.2 − cfa40 8.7 359.0 276.6 − − 24.8 − − bfa40 8.7 361.4 278.4 − − − 20.3 − cfa50 7.3 357.3 274.9 − − 31.0 − − cfa50b 7.3 385.3 296.9 − 2.9 31.0 − − bfa50 7.3 360.3 277.6 − − − 25.4 − cfa60 5.8 355.6 274.0 − − 37.2 − − bfa60 5.8 359.1 276.7 − − − 30.5 cfa49.5hl0.5 7.3 387.1 283.6 − 1.7 30.7 − 0.2 cfa49.5bfa0.5 7.3 387.1 293.4 − 3.8 30.7 0.3 − cfa48.8bfa1.3 7.3 387.5 311.0 − 3.8 30.2 0.6 − cfa45bfa5 7.3 387.3 309.6 − 1.3 28.3 2.6 − cfa49.5bfa0.6hl0.6 7.3 393.8 307.3 − 0.8 30.2 0.3 0.3 pc portland cement; sp superplasticizer table 2. binder inventory results of the material and transportation inputs for each concrete (per m3 of produced concrete) [4]. of the country), it is possible to known how many tons of coal are consumed to produce 1 kwh of electricity as well as the quantity of cfa produced during coal combustion. in portugal, the commercial value of cfa is about 18e/ton and the value of the electricity is 0.16e/kwh. therefore, the economic allocation coefficient of 0.17 % is applied to the impacts of the extraction, transportation and combustion of the coal from that power plant. as with the cement plant, this environmental report only covered the impact categories that are mandatory according to portuguese environmental legislation. as a result, all the flows (inputs and outputs) declared in this report were introduced in the simapro software, taking into consideration the quantified economic allocation coefficient of 0.17 %. regarding bfa, it is important to highlight that in portugal this kind of fly ash is considered a waste product and therefore they do not provide an economic value. because of this fact and according to the allocation rules presented in iso 14040, no flows from the thermal power plant are allocated in the production of bfa. with respect to the life-cycle inventory of the other used materials (gravel, sand, water and superplasticizer), generic data was used. since the development of specific environmental information for products is very time and cost consuming, initial lca studies, whose main goal was to compare different design scenarios, are normally based on generic (average) data [4,6]. for this reason and due to the lack of public available specific data for the abovementioned materials, this information was gathered from one of the most internationally accredited generic environmental databases of the ecoinvent report v2.2 [11]. the nearest context to the portuguese one was considered for this study. since the energy consumed during the manufacturing process is the parameter that most influences the life-cycle environmental impact [12] and 612 vol. 33/2022 biomass fly ash on the concrete sustainability since the portuguese energy mix is different from the european average [13], a contextualization of the energy used in each process was developed. this means that all used processes from the ecoinvent database were edited, and all energy input flows were changed to consider the portuguese energy mix. in the inventory of the transportation processes, the study considered the distances between the portuguese places of raw material extraction or raw materials storage facilities and the concrete mixing plant in question [4]. the distance between the raw materials production facilities and concrete production was considered for this study. the kilometres considered as the transportation distance of cement (secil group), sand and gravel (mibal minas de barqueiros, s.a.), coal fly ashes (pegop-energia eléctrica sa), biomass fly ash (altri), superplasticizer (basf) and hydrated lime (lusical-companhia lusitana de cal sa) to the concrete industry considered in this study were 41.5, 369, 177, 145, 329 and 117 km, respectively. the inventory related to the production of concrete was quantified taking into account the environmental product declaration (epd) of a specific portuguese concrete plant [14], where the different concrete formulations are supposedly produced. from this epd, only the flows related to the concrete mixing phase were considered. 2.4. impact assessment the life cycle inventory data was converted into potential environmental impact, using the lifecycle impact assessment (lcia) methods. in mars-sc, the definition of the sustainability indicators depends, above all, on the type of analysed product or building element and on the aims of the study. in this method, the environmental performance assessment is based on the following six environmental impact categories: i) global warming; ii) ozone depletion; iii) acidification of soil and water; iv) eutrophication; v) photochemical ozone creation; and vi) depletion of abiotic resources-fossil fuels. compared with the list of impact categories found in the en15804:2012 [15] standard, mars sc does not consider the depletion of abiotic resources-elements as an impact category. 2.5. normalization, aggregation and global assessment to avoid the scale effects in the aggregation of parameters of the different indicators and in order to minimize the possibility that some of the parameters are of the type. higher is better and lower is bad, it was necessary to normalize the indicators [15]. the normalization was done using the diaz-balteiro [16]. after normalization, is important to calculate the aggregation of each environmental indicator in terms of a global indicator, describing the overall environmental performance (nda). the global indicator nda results from the weighting average of each normalized. the sum of all weights must be equal to 1. for aggregation purposes, this study considers the weights (table 3) set in a study developed by the us environmental protection agency’s science advisory board (sab) [17]. in each sustainable profile, the global performance of a respective concrete with fly ash was monitored and compared with that of the reference concrete. indicator weight (%) gwp 38 odp 12 ap 12 ep 12 pocp 14 adp_ff 12 table 3. weight for each environmental indicator [17]. 3. results and discussion table 4 summarizes the values obtained in the quantification of the environmental impact categories, related with the production of the different concrete mixes. analysing the results, when it used just cement as binder, concrete presents the highest values of co2 emission. the high emission of co2 is a result of the chemical reactions (calcination) that occur during clinker production [4,18]. the incorporation of the bfa blended with cfa allows a reduction in all environmental impacts, moreover the environmental impacts decrease with increasing of cement substitution. at this stage, it is necessary to highlight the effect of the allocation step on the obtained results. in portugal, bfa are considered a waste product without economic value [18] and therefore there are no flows from the biomass power plant allocated to its production [4]. the same does not happen with cfa, as they have a market value and consequently a percentage of the power plant’s flows is allocated to their production [19]. concrete mixes 0.5%wt and 1.3%wt of bfa decreased the environmental impact related with the co2 emissions but an increase on the values of the others environmental indicators was observed [4]. this is a not result of the incorporation of bfa, since this result is not noted in the mixes with 5, 20, 40 and 60%wt, but it is due to the fact of these two mixes presented a higher superplasticizer content. it was concluded in other studies that superplasticizer had high influence on e.g. odp and adp_ff impact categories, but did not have a significant influence on gwp that most influences the overall environmental impact concrete with ashes, which confirm the results achieved in this study [4]. the normalization of the values obtained for each environmental impact category was obtained and the 613 e. r. teixeira, a. camões, f. g. branco acta polytechnica ctu proceedings formulations gwp100 odp ap ep pocp adp_ff kg (×102) kg (×10−5) kg kg (×10−1) kg (×10−2) kg (×103) ref 4.02 2.67 1.00 2.53 4.00 2.71 cfa20 3.44 2.55 0.91 2.32 3.64 2.56 bfa20 3.44 2.55 0.91 2.27 3.64 2.53 cfa40 2.83 2.38 0.80 2.08 3.23 2.36 bfa40 2.83 2.38 0.80 1.98 3.22 2.30 cfa50 2.53 2.30 0.75 1.96 3.02 2.26 cfa50b 2.88 2.69 0.90 2.41 3.97 3.10 bfa50 2.53 2.29 0.75 1.83 3.01 2.18 cfa60 2.23 2.21 0.69 1.84 2.82 2.16 bfa60 2.23 2.21 0.69 1.69 2.80 2.07 cfa49.5hl0.5 2.75 2.53 0.84 2.22 3.55 2.77 cfa49.5bfa0.5 2.96 2.77 0.94 2.51 4.20 3.31 cfa48.8bfa1.3 2.99 2.85 0.95 2.54 4.26 3.36 cfa45bfa5 2.75 2.59 0.85 2.22 3.57 2.73 cfa48.8bfa0.6hl0.6 2.73 2.57 0.83 2.18 3.46 2.64 table 4. values obtained for the different environmental impacts. formulations gwp100 odp ap ep pocp adp_ff [kg co2 eq] [kg cfc-11 eq] [kg so2 eq] [kg po4 eq] [kg c2h4 eq] [mj eq] ref 0.00 0.24 0.00 0.01 0.18 0.50 cfa20 0.32 0.44 0.30 0.26 0.42 0.62 bfa20 0.32 0.44 0.31 0.32 0.43 0.65 cfa40 0.66 0.72 0.65 0.54 0.71 0.78 bfa40 0.66 0.72 0.65 0.66 0.71 0.82 cfa50 0.83 0.86 0.82 0.68 0.85 0.85 cfa50b 0.64 0.20 0.32 0.16 0.20 0.20 bfa50 0.83 0.86 0.83 0.83 0.86 0.91 cfa60 1.00 0.99 0.99 0.82 0.99 0.93 bfa60 1.00 1.00 1.00 1.00 1.00 1.00 cfa49.5hl0.5 0.71 0.47 0.52 0.37 0.49 0.46 cfa49.5bfa0.5 0.59 0.09 0.21 0.04 0.04 0.03 cfa48.8bfa1.3 0.58 0.00 0.17 0.00 0.00 0.00 cfa45bfa5 0.71 0.37 0.50 0.38 0.47 0.49 cfa48.8bfa0.6hl0.6 0.72 0.41 0.54 0.42 0.55 0.56 table 5. normalized values of the studied environmental impact categories [4]. results are presented in table 5. this enables a better perception of which of the concretes has a better environmental performance. it is observed that, concrete in which 60% bfa had the best environmental performance, being ref and cfa48.8bfa1.3 the mixes that presented the lowest environmental performance. figure 1 present the sustainability profiles and the overall environmental performances. at the level of each impact category, the best concrete is the one that has the value closest to one. it is verified that the bfa60 concrete presented the best environmental performance (nda = 1.00) and plain cement concrete (ref) presented the worst performance (nda = 0.12) [4]. therefore, these results allow the conclusion that using a high content of bfa significantly increases the environmental performance of concrete production, since the overall environmental performance of concrete is improved. additionally, the usage of these materials contributes to a better compatibility between the construction sector and the goals of sustainable development [4]. 4. conclusions bfa used alone or blended displayed a capability to reduce the environmental impacts of concrete, when compared to conventional concrete. the results showed that 60% of bfa presented the best results. besides the good results observed, it is important to refer that the concrete formulations with 0.5 and 1.3%wt of bfa presented a low environmental performance, but better than that of plain cement concrete. the environmental issues related with 614 vol. 33/2022 biomass fly ash on the concrete sustainability figure 1. normalized values that described the sustainability profile. these two concrete formulations are due to the use of a higher content of superplasticizer when compared with the other formulations. the effect of superplasticizer on the environmental indicators is known, but the results showed that it is possible to slightly reduce the content of sp, without having a significant effect on the workability, maintaining the values similar to those of the concrete formulation with 5%wt of bfa [4]. the decrease of the content of sp has an important contribution on the improvement of environmental performance of these two concrete mixes. acknowledgements the authors wish to thank the portuguese foundation for science and technology (fct) and the eco-construction and rehabilitation doctoral program for supporting the phd scholarship (reference pd/bd/ 52661/2014). this work was partly financed by fct/mctes through national funds (piddac) under the r&d unit institute for sustainability and innovation in structural engineering (isise), under reference uidb/04029/2020. references [1] r. rajamma, l. senff, m. j. ribeiro, et al. biomass fly ash effect on fresh and hardened state properties of cement based materials. composites part b: engineering 77:1-9, 2015. https: //doi.org/10.1016/j.compositesb.2015.03.019. 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[19] d. garcía-gusano, d. garraín, i. herrera, et al. life cycle assessment of applying co2 post-combustion capture to the spanish cement production. journal of cleaner production 104:328-38, 2015. https://doi.org/10.1016/j.jclepro.2013.11.056. 616 https://repositorium.sdum.uminho.pt/bitstream/1822/7542/1/artigo%204.pdf https://doi.org/10.1016/j.ecolecon.2004.02.005 https://doi.org/10.1016/j.jclepro.2015.09.124 https://doi.org/10.1016/j.jclepro.2013.11.056 acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0034 acta polytechnica ctu proceedings 26:34–38, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app experimental and numerical investigation of cement paste microstructure and its hydration under short-term exposure to acetone yuliia khmurovska∗, petr štemberk, jiří němeček, magdaléna doleželová czech technical university in prague, faculty of civil engineering, department of concrete structures, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: yuliia.khmurovska@fsv.cvut.cz abstract. the solvent exchange methods are commonly used in order to arrest cement hydration reaction. this paper presents preliminary results of experimental investigation of cement paste microstructure under short-term (24 hours) exposure to acetone as a solvent in order to estimate the influence of solvent soaking time. the methodology to determine the effect of soaking time based on numerical prediction is also presented and described. the immersion of cement samples with the cross-section of 10 × 10 mm in acetone for 24 hours at the sample age of 15 hours does not enable to fully stop the hydration reaction, however, it can slow down the hydration reaction significantly. according to the comparison of several measured data (nanoindentation, scanning electron microscopy, observation made by optical microscope, mercury intrusion porosimetry) at the sample age of 1 month with the numerical simulation, the equivalent sample age is determined as approximately equal to 69 hours. keywords: acetone, cement paste, experiment, hydration reaction, numerical simulation. 1. introduction the hydration process of cement paste often has to be stopped in order to allow precise investigation and measurements of the alteration of cement paste properties over time. there are a numerous amount of methods to stop the hydration reaction [1–4], which can be divided into the following two groups: (1.) direct drying. the essence of the direct drying methods is to transform water into the vapor. the water vapor may escape freely and hydration reaction stops due to the lack of water, however, the direct drying methods are accompanied by the rapid change in temperature or pressure, which may damage microstructure of the sample or alter the cement paste properties and distort the results of the measurements [5–7]. the following types of direct drying are used in the available literature: • oven drying [5, 8, 9], • microwave drying [10–12], • vacuum drying [13], • freeze drying [14], • d-drying [15], • p-drying [15], etc. (2.) solvent exchange. the mechanism of solvent exchange methods is following: the solvent penetrates into the sample, mixes with free water and evaporates rapidly due to its low boiling point. it is believed that solvent exchange methods have lower impact on cement paste microstructure, however, it was reported [16, 17] that solvents may react with cement paste hydrates and alter its properties. also, it is difficult to entirely remove solvents, therefore, the chemical composition of the cement paste may be affected by the solvent. the following types of solvents are used in the available literature: • acetone [8, 18], • methanol [19, 20], • ethanol [21, 22], • isopropanol [14, 19], • tetrahydrofuran [1], • dimethyl sulfoxid [20], • benzene [16], • pentane [23], etc. from the above, it is obvious that the method to stop the hydration reaction should be chosen according to the subject of interest. this study is focused on cement paste microstructure investigation, therefore, the solvent exchange method was chosen and the acetone was used as the solvent. the necessary solvent soaking time to stop the hydration reaction depends on many factors, such as porosity, diffusivity, sample cross-section. porosity and diffusivity of early age cement paste change rapidly over time, therefore, the necessary solvent soaking time has to be a function of cement paste age. the majority of the research does not provide a sufficient explanation of the chosen soaking time. 34 https://doi.org/10.14311/app.2020.26.0034 https://ojs.cvut.cz/ojs/index.php/app vol. 26/2020 another complication is that it is impossible to determine whether the reaction was stopped during the solvent soaking and the additional experiments are needed for this purpose. therefore, the experimental investigation of the influence of solvents on cement paste has to be performed in order to determine the necessary solvent soaking time to stop the hydration reaction of early-age cement paste. this paper presents preliminary results of the experimental investigation of early-age cement paste microstructure change under short-term (24 hours) exposure to acetone in order to estimate the influence of acetone soaking time. the methodology to determine the effect of soaking time based on the numerical prediction is also presented and described in detail below. 2. experiment the cement paste which consists of cem i 42.5r extracem and water with water-to-cement ratio of 0.38 was used to produce small-scale 10x10x80 mm samples. the cement clinker composition is shown in table 1. the samples together with the plexiglass moulds were insulated by polyethylene foil till the age of 15 hours. after demoulding, all the samples were immersed in acetone for 24 hours at the sample age of 15 hours in order to stop the hydration process. after that all the samples were dried in oven at 50 °c till equilibrium in order to remove acetone. the samples were packed in polyethylene foil and were stored in box with silica gel until the examination in order to avoid possible moisture ingress [24]. environmental temperature and relative humidity were recorded starting from the mixing till packing with the interval of 1 minute. the average environmental temperature was 20.7 °c and relative humidity was 72.8% (excluding oven drying). the amount of anhydrous clinker was obtained using three different techniques, namely, nanoindentation (nano), scanning electron microscopy (sem) and observation made by optical microscope (opt) at the sample age of one month. the surface of the sample sections, which were used for nanoindentation and sem and opt investigation, was polished with silica carbide paper to achieve a smooth surface. the nanoindentation of the polished sample sections was carried out with a hysitron tribolab ti-700 with the berkovich diamond tip. five grids of 10 x 10 indents with the spacing of 10 µm were predetermined on the center part of the sample surface. the nanoindentation was performed using the force control test with the trapezoidal loading diagram. the sample surface was loaded by the diamond tip linearly during three seconds till the maximum force of 2 mn was reached with a following 20 seconds holding period and 3 seconds of unloading. the indentation elastic modulus was determined from the obtained load-displacement curves using the oliver and pharr theory [25]. the values of the elastic modulus higher than 50 gpa and hardness higher than 1.5 gpa correspond to the anhydrous clinker. after the nanoindentation, the polished sample sections were dried in a vacuum for three hours with a vacuum pump. the microstructure investigation of these sample sections was carried out by a phenom xl desktop sem and optical microscope. the amount of anhydrous clinker was obtained using the image processing technique which was performed by the fiji code. the measured volume fractions of anhydrous clinker and the average value are shown in table 2. the sample porosity and its distribution was measured using mercury intrusion porosimetry (mip) at the sample age of one month. the sample sections were broken into pieces and the center pieces were selected for further porosity investigation by the mercury porosimeter pascal 140 a 440. the measured porosity was in the range from 32.6% to 36.7% for different samples. more detailed description of the experiment is given in [24]. 3. numerical simulation in order to understated whether the hydration reaction was fully stopped after the samples immersion in acetone, the matlab code of hydration and microstructure models, which is based on the models described in [26–28], was developed. the data presented in table 1 and the measured environmental conditions were used in the analysis. the preliminary results are described below. the degree of hydration was calculated using the multi-component model, where the rate of cement hydration heat generation is obtained from the sum of the hydration heat rates of cement components according to their proportions in cement. the degree of hydration was calculated according the following formula αhyd = ∑ pi · (qi/qi,∞), (1) where pi is the mass ratio of the chemical component i of cement (c2s, c3s, c3a, c4af, gypsum, etc.), qi is the accumulated heat generation of the chemical component i, qi,∞ is the maximum theoretical specific heat of the component i and αhyd is the averaged degree of hydration. the calculated degree of hydration of the individual clinker components and their average and time when the samples were immersed in acetone are shown in figure 1 [24]. the cement paste microstructure was modeled using the expanding cluster model, volumetric balance and simplified calculation of the total pore surface area described in detail in [26–28]. after that the pore size distribution was calculated using a simplistic raleighritz distribution function. 35 y. khmurovska, p. štemberk, j. němeček, m. doleželová acta polytechnica ctu proceedings c3s, [%] c2s, [%] c3a, [%] c4af, [%] 63.99 14.30 9.11 8.26 c3s: alite; c2s: belite; c3a: celite; c4af: felite. table 1. clinker composition of cem i 42.5r extracem according to the manufacturer [24]. nano, [%] sem, [%] opt, [%] average, [%] 18.40 18.30 18.89 18.53 table 2. volume fractions of anhydrous clinker. figure 1. calculated degree of hydration [24]. figure 2. amount of anhydrous clinker. the volume fraction of anhydrous clinker was calculated as follows vcl = (1 − αhyd) · wp/ρp, (2) where wp is the weight of powder materials per unit paste volume, ρp is the density of powder materials and vcl is the volume fraction of anhydrous clinker. the decrease of the amount of anhydrous clinker due to the hydration reaction over time with the depiction of the averaged measured value is shown in figure 2 [24]. the capillary porosity was calculated using the following equation φc = 1 − vs − vcl, (3) figure 3. capillary porosity. figure 4. pore size distribution. where vs is the volume of hydrated products and vcl is the volume fraction of anhydrous clinker. the pore size distribution was calculated as follows φ(r) = φl + φg · vg + φc · vc, (4) where φl, φg and φc is the interlayer porosity, the gel porosity and the capillary porosity, respectively, vi = 1 − exp(−bir) is the normalized pore volume, bi is the porosity distribution parameter and r is the pore radius. the calculated capillary porosity over time with the depiction of measured porosity (the total porosity 36 vol. 26/2020 obtained using mip is considered to be predominantly capillary porosity) is shown in figure 3, while the measured and calculated pore distribution are shown in figure 4 [24]. as can be seen in figure 2, figure 3 and figure 4, the hydration reaction was not fully stopped at the sample age of 15 hours after sample immersion in acetone for 24 hours, but was slowed down significantly and the equivalent sample age is approximately equal to 69 hours according to the preliminary results of the numerical simulation at the real sample age of 1 month [24]. however, more detailed examination of the sample microstructure is still needed. 4. conclusions the immersion of cement samples with the crosssection of 10 x 10 mm in acetone for 24 hours at the sample age of 15 hours could not fully stop the hydration reaction, however, it can slow down the hydration reaction significantly. according to the comparison of the measured data at the sample age of 1 month with the numerical simulation, the equivalent sample age is approximately equal to 69 hours. the calculated volume fraction of anhydrous clinker, capillary porosity and pore size distribution have a good correlation with the measured values. however, the additional experiments with the variable acetone soaking time and different sample cross-section are necessary in order to draw the final conclusions. list of symbols αhyd averaged degree of hydration [–] ρp density of powder materials [kg m−3] φc capillary porosity [–] φg gel porosity [–] φl interlayer porosity [–] bi porosity distribution parameter [–] qi accumulated heat generation of chemical component (c3s, c2s, c3a, c4af) [kcal] qi,∞ maximum theoretical specific heat of component (c3s, c2s, c3a, c4af) [kcal] vcl volume of anhydrous clinker [m3 m−3] vs volume of hydrated products [m3 m−3] vi normalized pore volume [m3 m−3] wp weight of powder materials per unit paste volume [kg m−3] pi mass ratio of chemical component of cement (c3s, c2s, c3a, c4af) [–] r pore radius [m] acknowledgements this work was supported by ministry of education, youth and sports of czech republic, project 8f17002, and the czech technical university, project sgs19/038/ohk1/1t/11, which are gratefully acknowledged. references [1] j. zhang, g. w. scherer. comparison of methods for arresting hydration of cement. cement and concrete research 41(10):1024–1036, 2011. 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https://doi.org/10.1201/9781482288599 acta polytechnica ctu proceedings 26:34–38, 2020 1 introduction 2 experiment 3 numerical simulation 4 conclusions list of symbols acknowledgements references https://doi.org/10.14311/app.2022.33.0039 acta polytechnica ctu proceedings 33:39–44, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague conformity assessment model for mechanical coupling devices and anchors for reinforcing steel for use in concrete structure lee brankley, ayhan turgul, lyn morgan, ladin camci∗ cares, pembroke house, 21 pembroke road, sevenoaks, kent tn13 1xr, united kingdom ∗ corresponding author: ladincamci@ukcares.com abstract. the use of mechanical coupling devices and anchors for reinforcing steel for use in concrete structures can simplify the design and construction of reinforced concrete and reduce the amount of reinforcement required. technical approvals of these products by suitably qualified organisations are acceptable under building regulations and by major specifiers as the appropriate way of assessing non-standard or new and innovative products. specifiers, manufacturers, traders, and customers increasingly demand certainty in the quality management of mechanical coupling devices and anchors entering the supply chain. the supply chain is extremely complex. the weak traceability systems and pressure on the supply chain to reduce cost have created the environment where risks increase. this unique and comprehensive model describes the technical and assessment requirements for the conformity of mechanical coupling devices and anchors. it covers all stages in the supply chain from the production, assembly, product performance and evaluation of mechanical properties including tensile strength, ductility, permanent elongation (slip), cyclic loading performance and fatigue performance. the essential elements of the quality management system combined with additional requirements on process control, product testing and robust installation methodology are essential and agreed by all relevant sectors of industry. keywords: certification, mechanical rebar anchors, mechanical rebar couplers, mechanical rebar splicing systems, product conformity. 1. introduction the cares technical approval scheme for couplers and anchors provides a comprehensive assessment process, ensuring complete product traceability to the construction site. a key part of assessment process is the development of a technical approval report which not only indicates the performance characteristics and conformity of the product, but also ensures that correct assembly to the rebar is undertaken with robust installation procedures contained within each report. the cares technical approval scheme predominantly covers couplers for use with bs 4449 [1] reinforcing steel. cares with its global presence have recognised the need for more global approvals and have recently introduced a scheme which not only supports cares approved rebar to bs 4449 [1], but also rebar to en 10080 [2], astm a615/a615m [3], astm a706/a706m [3], iso 6935-2 [4] and ss 560 [5]. the cares technical approval certification scheme for the supply and installation of couplers and anchors has been developed in conjunction with the cares technical committee, made up of various interested parties including leading structural engineers, specifiers, and coupler producers. the schemes are based on the demonstration of the continuing operation of the quality management system consistent with iso 9001 [6], the appropriate quality and operations assessment schedule [7] and the ability to manufacture, supply and / or install material or product complying with bs 8597 [8], sellafield engineering standard [9], highways england manual [10] or more recently, iso 15835 [11]. at the initial approval stage, all quality affecting manufacturing processes are audited whether produced internally or outsourced via approved subcontract suppliers. once approved, all cares approved coupler / anchor manufacturers must ensure any design, material specification, supply chain or installation instruction changes are vetted by cares to ensure any necessary testing is further carried out to ensure product conformity is maintained prior to the change. the cares coupler technical approval scheme is further supported both downstream and upstream by the cares steel certification scheme where reinforcing bar fabricators must also be cares approved for installation of couplers on the reinforcing steel (cares src08 scheme). 2. objective the objective of the technical approval scheme is to give confidence to the purchaser that the construction 39 https://doi.org/10.14311/app.2022.33.0039 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en l. brankley, a. turgul, l. morgan, l. camci acta polytechnica ctu proceedings schedules no. description ta1-a technical approval of couplers for reinforcing steel for use in structures and structuralelements designed in accordance with the fatigue requirements of structural eurocodes ta1-b technical approval of couplers for reinforcing steel and reinforcement anchors for bs 8110 [12] and en 1992-1-1 [13] applications for static loading in tension or tension and compression. ta1-c technical approval of tension couplers for reinforcing steel for sellafield engineeringstandard [9] applications ta1-f technical approval of couplers for high cycle fatigue and low cycle loading and static loadingapplications in tension. src08 quality and operations assessment schedule [7] for the application of mechanical couplersto reinforcing steel table 1. quality and operations assessment schedules [7]. product complies with the appropriate specification without the need to undertake separate verification. additionally, it provides confidence in both traceability and correct installation methodology. this involves the application of quality assurance principles to assess the supplier’s and supply chain quality systems and, as appropriate, product testing to ensure conformity with the requirements contained in each scheme at each part of the supply chain. 3. assessment and certification system 3.1. principles the assessment and certification system employed in the scheme has been developed around the following common principles: • the scheme is concerned with ensuring that the construction product complies with the relevant product standards and / or client specifications. • the responsibility for complying with the standards rests absolutely with the approved firm. • the means of providing assurance of consistent compliance with the standards or specifications shall be the formal management system for quality, including process control, which the approved firm is required to have documented and to be operating to the satisfaction of cares. this system is subject to ongoing audits at periodic intervals. • a key method of verifying compliance of the product is by the independent assessment and surveillance of the approved firms management system for quality, including witness testing of the product as appropriate • additionally, independent inspection and testing of the product is undertaken by testing laboratories approved by cares. • long term quality level reviews and quality complaints are a key feature of the ongoing surveillance programme. 3.2. quality and operations assessment (technical assessment) schedules the schemes quality and operations assessment schedules [7] for couplers and anchors are shown in table 1. 3.3. assessment the assessment consists of: • full assessment in accordance with the relevant quality and operations assessment schedule [7]. • the manufacturing process route integrity, including records, resources, training, contract review, product requirements, manufacturing controls and interfaces. • review of installation methodology and product dimensional conformance. • product conformity assessment by sampling in accordance with the relevant quality and operations assessment schedule [7]. • preparation and issue of the cares technical approval report for the couplers / anchors assessed. 3.4. applicant the applicant is required to have and be able to demonstrate: • the ability and resources required to produce the coupler / anchor to the relevant standard(s) and / or specifications consistently. • a quality management system which complies with the requirements of iso 9001 [6], and the relevant quality and operations assessment schedule [7]. the system shall be described formally in the supplier’s quality manual. • adequate control of approved suppliers and incoming product. • the ability of the coupler manufacturer to control its supplier is essential in the case of approval of cementitious grouted couplers, where the grout has 40 vol. 33/2022 model assessment of reinforcing steel in concrete structure tight performance requirements as part of the latest coupler technical approval scheme ta1-f. • robust installation methodology which is readily transferrable to downstream fabricators or site installers. 3.5. certification a recommendation for approval is produced when the manufacturing procedures and product conformity have been assessed and the scheme criteria have been satisfied and all reported deficiencies have been addressed to the satisfaction of cares. a key part of this recommendation process is the preparation and subsequent approval of the coupler / anchor technical approval report. copies of all the issued technical approval reports are freely and readily available to download from the cares website (www.ukcares.com). the certificate of approval is valid for one year with renewal subject to continuing satisfactory performance of the construction product. firms granted certificates of approval are required to reproduce the cares logo, and the certificate of approval number on all relevant documentation. 3.6. ongoing surveillance the assessment and certification system employed in the scheme has been developed around the following common principles: firms granted certificates of approval are required to undergo surveillance audits by cares twice per annum, to ensure continued compliance with the required scheme and product standard. samples of the construction product may be witnessed tested at the manufacturer in addition to the requirement to test "mirror" samples at an independent laboratory approved by cares. all technical approval schemes require that key variables of the original tested component remain unchanged. the following variables as applicable to the type of approved coupler shall not be changed without the prior agreement and subsequent assessment by cares: a) coupler component material or grade or process route of the raw material used to manufacture the coupler b)cross-sectional area of the splice sleeve c) bar engagement length d)reinforcing bar profile (rib pattern), grade or specification e) bar locking device f) thread geometry g) torque h)swaging pressure i) die geometry j) production process or production location k) approved grout supplier and / or change in grout specification or properties (in the case of cementitious grouted coupler) l) installation instructions a key condition of the cares technical approval schemes are that any subsequent changes to the above must be reviewed by cares prior to the implementation of the change, thus ensuring that any extra testing is undertaken to ensure continued product conformity. failure to do so will usually lead to the removal of the technical approval and the cares approval certificate. new suppliers deemed critical for the processed part will be assessed by cares before production can commence. all sub-suppliers are visited at least once every three years as part of the ongoing surveillance installation instructions are validated at each surveillance visit, by means of witness sampling of the coupler and application to the reinforcing bar. 3.7. key attributes of the approved coupler / anchor the approved coupler / anchor attributes include a) tensile properties, ensuring the connected reinforcing steel properties are not compromised b)slip < 0.10 mm (permanent elongation after unloading from 60% / 65% of the characteristic yield strength of the reinforcing steel). this is key requirement in all the cares coupler schemes not only at the initial assessment but also ongoing factory production control surveillances. control on slip is essential to ensure adequate concrete crack control. c) mode of failure, within or outside the coupler assembly. d)performance in low cycle / high cycle fatigue according to each cares scheme appropriate to the coupler final application, whether highway bridges, critical installations such as nuclear or areas of potential seismic activity. e) designed to allow consistent and robust installation methods readily repeatable whether at fabricators or on site 3.8. methodology of conformity assessment and ongoing surveillance for the coupler / anchor approval key requirements of the cares assessment and surveillance methodologies are: a) use of assessors who have a background in the steel production and / or construction industry is critical. these assessors will also have a thorough knowledge of principals of iso9001, with clear 41 www.ukcares.com l. brankley, a. turgul, l. morgan, l. camci acta polytechnica ctu proceedings attention to factory production control (fpc), maintenance of traceability, construction product properties and installation methodology. b)the manufacturing site visits, including visiting of any key sub-contract operations are agreed with the client in advance and an audit itinerary issued to ensure full understanding of the requirements. c) a key part of each audit of the client and any subcontractors is to establish confidence in their fpc and associated procedures, purchase control and specifications, incoming inspection of goods and final product inspection and release. d)management of change is facilitated using schedules of operations which are required to be produced by each approved client and required to be updated regularly to detail any changes to the agreed manufacturing process route, critical subcontractors, purchasing specifications and product drawings. also included are the issue status of operational procedures for any rebar end preparation / threading machines leased to cares approved fabricators. these schedules are reviewed during audits to ensure there an adequate management of change record, thus ensuring confidence in the maintenance of the product properties. e) sampling of the assembled product is a key feature at each audit whether at the manufacturer or fabricator following the agreed installation procedure. these are subsequently tested at independent laboratories which are approved by cares f) continued approval and certification is only allowed following successful auditing and product testing. any areas for improvement are followed up the issue of non-conformances which require the client to undertake a root cause analysis, immediate remediation, and corrective actions. 4. identification and traceability system cares certification provides full traceability from the molten steel to the construction site by addressing all stages of production, delivery, and receipt. once installed and cast into concrete within the structure, construction components cannot be easily removed even if there is a product failure or a manufacturing deficiency. there is no ’product recall’ when things go wrong with steel in the structure of a prized project. its performance must be assured from the outset. cares have embraced the digitalisation era by developing and promoting its cares cloud. this cares digital platform plans to cover all stages in the supply chain from receipt of raw materials, the manufacture and processing of steel products through to delivery to the end user. it delivers product test results and sustainability credentials that are traceable across the whole supply chain ensuring that rebar is effectively tested, marketed, and traced. critically, in the latter stages of this development it could also confirm where these products are installed in the structure. 5. construction product supply chain cares approach the supply chain is designed to give confidence to the purchasers and the technical approval schemes are supported by the steel schemes as shown in figure 1. for a product to be cares approved, all the supply chain must be approved, whether it is the steelmaker, roller, steel trader or fabricator, there shall be no break in this link. coupler installation is rigorously reviewed during surveillance at approved fabricators as part of the cares steel scheme src08, ensuring agreed installation procedures are maintained in line with those detailed in the appropriate technical approval report. witnessed coupled rebar is taken for subsequent independent testing at an agreed rate over the surveillance period. 6. assessment models and specification • cares ta1-a quality and operations schedule for the technical approval of couplers for reinforcing steel for use in structures and structural elements designed in accordance with the fatigue requirements of structural eurocodes: this scheme is based on the requirements of the highways england [10] specification for highways structures, with continual product assessment by independent testing and ltql. testing includes both tensile and high cycle fatigue testing. • cares ta1-b quality and operations schedule for the technical approval of couplers for reinforcing steel and reinforcement anchors for static loading in tension or tension and compression: this scheme is based on the requirements of bs 8597 [8], with continual product assessment by independent testing and ltql. testing includes both tensile and optional compression testing. • cares ta1-c quality and operations schedule for the technical approval of tension or tensioncompression couplers for reinforcing steel and reinforcement anchors for nuclear applications based on sel lafield engineering standard [9]: this scheme is based on the requirements of the sellafield engineering standard 42 vol. 33/2022 model assessment of reinforcing steel in concrete structure figure 1. concept of ’extended’ product responsibility applied to reinforcing steel. [9], with continual product assessment by independent testing. testing includes both tensile, including cold soak tensile testing, low cycle, and high cycle fatigue testing. • cares ta1-f quality and operations schedule for the technical approval of couplers for high cycle fatigue and low cycle loading and static loading applications in tension: this scheme is based on the requirements of iso 15835 [11] with enhancements. continual product assessment in line with the standard by independent testing. testing includes tensile, optional low cycle and high cycle fatigue testing. this is the first cares scheme to include a cementitious grouted coupler, hence evidence of conformance and product performance of the specified grout is a key part of this scheme. this built on cares experience of using grout on site for bonded tendons in post tensioning installations. 7. conclusion cares is an independent, not-for-profit certification body. the certification scheme for both the steel supply chain and installation of couplers in its requirements, providing for good management control, consistent product and technical conformity, industry developed disciplines, rules and standards and auditing by industry professionals with many yearsexperiences in the appropriate industry. cares reduce the risk of using construction products which do not meet the specification requirements, and to enable feedback and corrective action in case problems arise. major construction clients, designers, and contractors worldwide, recognise and even specify cares and its certification scheme to gain assurance of consistent product quality and technical competence. cares maintain its construction products development by attendance at the european organisation of technical approval (eota) meetings and actively participate in various iso, en and bs standard meetings. cares schemes are prepared in collaboration with the relevant technical committees of interested parties which include senior structural engineers, specifiers, and producers. references [1] bs 4449:2005+a3:2016: steel for the reinforcement of concrete bar, coil and decoiled product, european standards, 2016. [2] en 10080:2005 steel for the reinforcement of concrete weldable reinforcing steel general, 2005. [3] astm a615/a615m: 2018 standard specification for deformed and plain carbon steel-bars for concrete reinforcement & astm a706/a706m:2016 standard specification for deformed and plain low-alloy steel bars for concrete reinforcement, 2018. [4] iso 6935-2:2019 steel for the reinforcement of concrete part 2: ribbed bars, 2019. [5] ss 560:2016 (singapore standard) specification for steel for the reinforcement of concrete weldable reinforcing steel bar, coil and decoiled product, 2016. [6] iso 9001:2015 quality management systems requirements, 2015. [7] cares quality and operations assessment schedules. www.ukcares.com. [8] bs 8597:2015 steels for the reinforcement of concrete reinforcement couplers requirements and test methods, 2015. 43 www.ukcares.com l. brankley, a. turgul, l. morgan, l. camci acta polytechnica ctu proceedings [9] sellafield engineering standard es_0_3110_2 issue-4 mechanical splices to reinforcement for concrete. part 2 manufacturing, installation and construction requirements. https://www.ukcares.com/tapdf/5067.pdf. [10] highways england mhwa amendment 2014 manual of contract documents for highways works volume 1, series 1700, 2014. [11] iso 15835-1:2018 steel for the reinforcement of concrete reinforcement couplers for mechanical splices of bars requirements, iso 15835-2:2018 steel for the reinforcement of concrete reinforcement couplers for mechanical splices of bars test methods & iso 15835-3:2018 steel for the reinforcement of concrete reinforcement couplers for mechanical splices of bars conformity assessment scheme, 2018. [12] bs 8110 structural use of concrete (replaced by eurocode 2), british standard. [13] en 1992-1-1:2004, eurocode 2: design of concrete structures part 1-1: general rules and rules for buildings, 1992. 44 https://www.ukcares.com/tapdf/5067.pdf https://doi.org/10.14311/app.2022.33.0571 acta polytechnica ctu proceedings 33:571–577, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague mechanical and durability performance of cementitious matrices with low cement and different additions eliana soldadoa, ∗, ana antunesa, hugo costaa, b, ricardo carmoa, b, eduardo júliob, c a isec-polytechnic institute of coimbra, department of civil engineering, rua pedro nunes quinta da nora, 3030-199 coimbra, portugal b ceris-ist, universidade de lisboa, av. rovisco pais, 1049-001 lisboa, portugal c instituto superior técnico universidade de lisboa, department of civil engineering, architecture and georesources, av. rovisco pais 1049-001 lisboa, portugal ∗ corresponding author: eliana.soldado@isec.pt abstract. nowadays, the decarbonization of the concrete sector is a priority to counteract the global environmental harmful. knowing that portland cement is responsible for significant anthropogenic emissions worldwide, the optimization of cement dosage on concrete mixtures is a key issue, combined with the mechanical and durability performances, allowing to improve its eco-efficiency. the goal of the present work consists in optimizing the formulation of concrete matrices, used to produce reinforced concrete poles for the electric lines, by increasing the packing density, reducing the cement dosage and incorporating different additions, in order to promote the performance. two different formulations of matrices were considered, fresh moulded concrete with dry consistency and formwork moulded concrete with plastic consistency. for each one, different dosages of cement and various additions (limestone filler, fly ash, natural pozzolan, electric furnace slag) were considered to replace the cement, which were used single or combined. the mortar matrices of the concrete formulation were characterised regarding: fresh properties; mechanical properties (flexural and compressive strengths and young’s modulus); shrinkage and durability properties (capillary water absorption, carbonation and chloride ingression resistances). the influence of the additions on those properties was analysed and compared with reference cement-based mixtures and the main conclusions are presented. keywords: additions, durability, formulation optimization, low cement concrete, mechanical performance. 1. introduction eco-efficient and economical concrete production has been achieved, among other strategies, by partially replacing portland cement with by-products from other industries. the great potential for the impact reduction of the environmental footprint related to cement production has been seen especially for concretes with current strengths, but also for high and ultra-high strengths. the granulometric optimization, combined with the use of cement with high reactivity and additions as cement substitutes, and the use of powerful superplasticizers together with the reduction of water dosage, allow a significant reduction of the clinker present in portland cement, which enables the production of concretes with low cement dosages and good mechanical performance [1, 2]. this evolution, from conventional mixtures to eco-efficient concrete, which is expected to be proved by studies of life cycle analysis, should not diminish the performance of cementitious materials and must take into account their workability, strength and durability performances. optimizing the formulation of this concrete has a potential positive influence on mechanical properties at longer ages due to pozzolanic effect, as well as also on time dependent and on durability properties. although the development of the initial strengths is slower compared to traditional concretes, it is possible to obtain concrete with higher mechanical strengths at older ages, because the pozzolanic reactions are slow but prolonged [2–4]. optimizing packing density is not only beneficial for reducing the water amount, but can also reduce the space between cement particles. to optimize the packing density of the concrete, these should be selected to fill the voids between large particles with smaller particles and so on to obtain a dense and stiff structure of concrete matrix, reducing the space that needs to be filled by hydration products, which results in higher strengths. the volume of water required to ensure workability can also be reduced by adding superplasticizers in the matrix [3–5]. the published works on this subject indicates that, in the production of eco-efficient concrete, a signifi571 https://doi.org/10.14311/app.2022.33.0571 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en e. soldado, a. antunes, h. costa et al. acta polytechnica ctu proceedings mixture binder cement flyash lim. filler pozz slag sand mix. coarse aggr. mix. w/b compact-ness (kg/m3) (kg/m3) (kg/m3) (kg/m3) (kg/m3) (kg/m3) (kg/m3) (kg/m3) (−) (−) bt ref 400 400 − − − − 930 902 0.390 0.824 bt200 fa 350 200 150 − − − 804 1144 0.312 0.865 bt250 350 250 50 50 − − 829 1129 0.323 0.860 bt200 poz 350 200 − − 150 − 797 1138 0.326 0.860 bt250 slag 350 250 − 50 − 50 861 1120 0.323 0.860 mt ref 360 360 − − − − 778 1046 0.472 0.808 mt250 fa 350 250 100 − − − 902 992 0.409 0.840 mt+200 fa 400 200 200 − − − 888 920 0.370 0.835 mt250 poz 350 250 − − 100 − 902 992 0.408 0.840 mt250 slag 350 250 − − − 100 957 982 0.408 0.840 table 1. formulation of concrete mixtures used to produce the mortar matrices. cant reduction in the amount of cement is achieved without changing the properties of the mixtures, thus reducing co2 emissions per cubic meter of concrete in 25% [3, 6]. filer and fly ash, for example, have been identified as viable additions with visible improvements in workability, mechanical properties and durability [1, 2]. fly ash is a by-product with moderate cost and an efficient addition to replace cement in concrete mixtures, with good results to assure strength and durability. however, its growing demand and uncertain availability in developed countries, also due to environmental concerns, demands alternative additions, such as natural pozzolans or other by-products, like slags. some studies point to more than 65% reduction in environmental impact when blast furnace slag is used for partial cement replacement [2]. industrial concrete, both in precast and ready-mixed production, is not optimized in terms of formulation or cost and efficiency. dosages of portland cement should be minimized to reduce production costs and to be more concerned with ecoefficiency, by including more additions in formulation. also, by maximizing the packing density and densification of the binder matrix, it makes the matrix more closed and provides greater durability and high mechanical strength with lower cement content. in this sense, mixtures in mortar matrix of the concrete, considering the partial replacement of cement by different additions such as limestone filer, fly ash, natural pozzolans or ground electric furnace slag, were studied in the present work, promoting sustainable concrete solutions. two types of mixtures were studied: one with dry consistency for fresh moulded concrete bt and another with plastic consistency concrete mt, for formwork moulding concrete. 2. materials and methods 2.1. materials to study the mixtures, the mortar matrix of the concrete was considered, consisting of binder paste and sand fractions. cem ii-a/l 42.5r (c) was used, as well as f-type fly ash (fa), limestone filler (lf), cape verde pozzolans (poz) and electric furnace slag (slag) additions, with the following density values in kg/gm3: 3.08; 2.30; 2.70; 2.30; 3.83, respectively. since sand has a great influence on the grain size of the mixtures, particularly on packing density and workability, a mixture of fine sand 0/1 mm and medium 0/4 mm sand was used. two types of admixtures were used: sikament he200p, a water reducer for dry concrete with density of 1.17 kg/gm3, favouring a better dispersion and utilization of cement for the bt dry mixtures; glenium sky526, ether-polycarboxylate based superplasticizer with a density of 1.06 kg/gm3, to increase plasticity and reduce the amount of water for the mt mixtures. the dosage set for each mixture was adjusted to achieve the desired target workability and the set air content. 2.1.1. mixtures the two types of mortar matrix mixtures, bt-dry consistency and mt-plastic consistency, were formulated according the proportioning of corresponding concretes, being those presented in table 1, together with water/binder ratio (w/b) and the compactness of each mixture, and compared with two reference mixtures, bt_ref and mt_ref, which have current design parameters of precast concrete. compactness is a packing density volumetric ratio between the volume of solid particles and the total volume of con572 vol. 33/2022 mechanical and durability performance of cementitious matrices figure 1. flexural strength at 28 and 56 days. crete, which is firstly estimated using the formulation parameters and then controlled by the measured values of density, air content and compacted volume. the air content was targeted to 2.0% in bt and 1.5% in mt mixtures, which corresponds to 3.5% and 2.4%, respectively in mortar matrices. 2.2. experimental tests the tests to measure the consistency through spreading and air content of the mixtures, in order to adjust the dosages of the constituents, were performed according to the procedures of en 1015-3 [7] and en 1015-7 [8], respectively. admixtures’ dosages were adjusted to achieve the target consistency of each mixture, whose dosage in relation to the weight of cement varied between 29 and 34.8 ‰ in bt and between 7.2 and 10.5 ‰ in mt. the tests to access flexural and compressive strengths on mortars were performed on prismatic specimens (40 × 40 × 160 mm3), according to en 1015-11 [9], and the results were obtained by averaging three specimens for flexural and six half specimens for compressive, at each age (28 and 56 days). shrinkage was assessed at several ages (1, 3, 7, 14, 28, 56, 90, 120 and 180 days) by the average of three prismatic specimens (40 × 40 × 160 mm3) according to en 1015-13 [10]. the young’s modulus test was performed after a 56-day of curing and followed the standard lnec e397 [11], using a prismatic specimen of 50 × 50 × 300 mm3. capillary water absorption was evaluated by averaging three specimens of 40 × 40 × 160 mm3 according to the lnec e393 [12]. the determination of carbonation resistance was assessed according nt build 357 [13] standard by averaging two specimens of 40 × 40 × 80 mm3 with 5% of co2 at different ages (7, 14, 28, 56 and 90 days). the chloride diffusion coefficient by non-steady state migration was obtained according to the procedure of nt build 492 [14], at 28 and 56 days, by the average of three specimens per age. figure 2. compressive strength at 28 and 56 days. 3. results and discussion 3.1. mechanical properties and shrinkage 3.1.1. flexural and compressive strengths the results of average strengths at 28 and 56 days, for bt and mt mixtures, are shown in figure 1 and 2, respectively for flexural and compressive strengths. both bt and mt formulations, regarding flexural strength, do not tend to evolve from 28 to 56 days, in which minor reductions were registered. despite the high reduction of cement, several bt mixtures present better performance than the reference, being that improvement even higher in mt. bt250 slag and mt 250 slag and mt+200 fa mixtures stands out, since they reach higher values than those of the references, with increases of 39%, 85% and 100%, respectively at 28 days, being those relations similar at 56 days. in compressive strength, for both formulations (bt and mt), the evolution from 28 to 56 days is significant, until 15%, due to pozzolanic effect. despite the cement reduction, mixtures present similar or higher performance than the references, by increasing packing density and incorporating additions. mixtures with fly ash and with slag in cement replacement revealed improvements until about 11%, and 6% respectively. 3.1.2. shrinkage evolutions of mortar shrinkage over time are shown in figure 3, for bt and mt mixtures. in the first type, bt250 mixture (with filler and fly ash additions) and bt200 slag stands out for its lower shrinkage results, being the shrinkage about 20 to 25% below reference; bt200poz mixture stands out for having a very high initial shrinkage, which stabilizes after the first month, being the value about 20% superior than reference; the other mixtures of this type present a smoother evolution, stabilizing after 90 days. in mt mixtures, all present lower values (from 20 to 50%) than reference, being mt250 fa and mt250 slag those with the lowest shrinkage; the evolution of those mixtures tend to stabilize between 28 and 56 days, while reference reaches that point at 120 days. 573 e. soldado, a. antunes, h. costa et al. acta polytechnica ctu proceedings � � ��� ��� ��� ��� ���� � �� �� �� �� ��� ��� ��� ��� ��� i̇ f v� �î �� � � w��gd\v� %7�uhi %7����)$ %7��� %7����3r] %7����6odj � ��� ��� ��� ��� ���� � �� �� �� �� ��� ��� ��� ��� ��� i̇ f v� �î �� � � w��gd\v� 07�uhi 07����)$ 07�����)$ 07����3r] 07����6odj figure 3. shrinkage evolution of bt (left) and mt (right) mixtures with time. ��� � ���� ���� ���� ���� ���� ���� ���� ���� ���� ���� � � �� �� �� �� �� �� �� %7������������������������������������07 ( ��* 3d � %7�uhi %7����)$ %7��� %7����3r] %7����6odj 07�uhi 07����)$ 07�����)$ 07����3r] 07����6odj figure 4. young’s modulus test (left) and the values for bt and mt mixtures at 56 days (right). 3.1.3. young’s modulus the obtained average values for the young’s modulus at 56 days are shown in figure 4. by analysing the graph, the additions and increasing of packing density is evident in increasing young’s modulus compared to reference, being that increase of 8 to 16% in bt mixtures and until 6% in mt mixtures. pozzolans, fly ash, and slag additions revealed significant increases, which may have an even higher effect on concrete, since high packing density and coarse aggregates will increase that property. 3.2. durability 3.2.1. water absorption through capillarity figure 5 shows the values of capillary water absorption of the bt and mt mixtures. since the concrete quality of this parameter is considered "high" below 0.1 mg/mm2 × min1/2, "average" between 0.1 and 0.2 mg/mm2 × min1/2 and "low" above 0.2 mg/mm2 × min1/2 [15], it is possible to verify that in bt mixtures, only the bt250 can be considered of medium quality, while the others are well rated. as for mt mixtures, only mt250 poz and mt250 slag remain throughout the test in the "high quality" classification. the rest start from being medium quality and acquire a better rating over time. this proves a nonlinear behaviour of this parameter, since the evolutions are more compatible with power functions, tending to improve with age by the pozzolanic effect. by analyzing the results of this test, performed on three consecutive days, it can be concluded that the mixtures with higher packing density, both bt and mt, combined with slag addition present the best results in the evaluation of this parameter, absorbing about 45% less water than the reference in bt mixture and less 60% in mt mixture. fly ash and pozzolan addition also revealed good performance in reducing water absorption. 3.2.2. carbonation resistance the analysis of carbonation depth, cdepth, of the studied mixtures (figure 6) proved to evaluated with a linear tendency with square root of time (figure 7), as expected. it is evident that the amount of cement is highly influent in carbonation resistance; despite the increase of packing density and the replacement of cement by additions, the optimized mixtures revealed in general higher values then the references, until 160% in bt200fa and 95% in mt+200fa. in bt mixtures, this phenomenon is more noticeable since the difference of cement dosage is higher and 574 vol. 33/2022 mechanical and durability performance of cementitious matrices � � � � � � � � �� �� �� �� �� 6d ��p j� p p � � w��plq���� 6 ��� 6 ��� %7�uhi %7����)$ %7��� %7����3r] %7����6odj � � � � � � � � �� �� �� �� �� 6d ��p j� p p � � w��plq���� 6 ��� 6 ��� 07�uhi 07����)$ 07�����)$ 07����3r] 07����6odj figure 5. capillarity water absorption through square root of time of bt (left) and mt(right) mixtures. ���gd\v������������������gd\v������������������gd\v� figure 6. chamber for the carbonation test and evolution of carbonation mt250fa with time. � � � � �� �� �� � � � � � � � � � & gh sw k �p p � w��gd\v���� %7�uhi %7����)$ %7��� %7����3r] %7����6odj � � � � �� �� �� � � � � � � � � � & gh sw k �p p � w��gd\v���� 07�uhi 07����)$ 07�����)$ 07����3r] 07����6odj figure 7. carbonation depth of bt (left) and mt (right) mixtures. � � � � � �� �� �� ���gd\v ���gd\v '��î�������p��v� %7�uhi %7����)$ %7��� %7����3r] %7����6odj 07�uhi 07����)$ 07�����)$ 07����3r] 07����6odj ���gd\v���������������������������gd\v� figure 8. chloride migration (left) and d coefficient at 28 and 56 days for bt and mt mixtures (right). 575 e. soldado, a. antunes, h. costa et al. acta polytechnica ctu proceedings the reference shows the best results, followed by mixtures with more cement, bt250 slag and bt250, although it contains partial substitution by additions. mt mixtures show the same trend, although with a smaller difference from the reference; in those, it is also possible to verify that the increase in the amount of binder replacing cement is harmful to the carbonation resistance, while mt250fa present good resistance, with a carbonation depth similar to the reference. mt250 slag present a significant reduction (about 30%) in comparison to the reference, proving the ability of slag to increase carbonation resistance. 3.2.3. chloride migration the results shown in figure 8 represent the chloride diffusion coefficient (d) by non-steady state migration test at 28 and 56 days. in this test (figure 8), as lower is d coefficient as higher is the resistance to chloride migration, being the durability also higher. the graph prove that mixtures improve its results at 56 days, comparing to 28 days, due to pozzolanic effect, in which in most cases this difference is substantial (reduction until near 65% for mixtures with fly ash). mixtures with pozzolan additions (bt200 poz and mt250 poz) have very low d results compared to the others and to the reference, which makes them stand out positively, being the coefficient until 85% and 80% lower than references, respectively. in contrast, mixtures with slag additions (both bt250 slag and mt250 slag) show the higher chloride migration, being until about 25% and 60% higher than reference in mt and bt mixtures, respectively; in slag addition does not promote apparent improvement with age, contrarily to the remaining mixtures. 4. conclusions from the analysis of the obtained results in the experimental study, regarding the effect of combining the increase of packing density with adding different cementitious supplementary materials in two types of mortar mixture (bt, with dry consistency; mt, with plastic consistency), it was possible to draw the following conclusions: 1. despite the reductions in cement content (from 30 to 50%), flexural strength increases significantly with different additions up to 100% comparing to reference mixtures, being the flay ash and slag mixtures those with higher strengths. 2. compressive strength of mixtures with additions tend to increase from 28 to 56 days up to 15% due to pozzolanic effect and, despite the cement reduction, it increases up to 6% and 11% with slag and fly ash, in comparison to references. 3. the replacement of cement by additions promotes a significant decrease in the shrinkage between 20 to 50%, being mixtures with slag and fly ash those with major reductions. 4. adding eco-efficient supplementary binders (pozzolan, slag and fly ash) increases young’s modulus of mortar matrices up to 16% (mainly in mixtures with dry consistency, bt), comparing to references, being expected to have a higher effect on corresponding concrete, due to the stiffness of coarse aggregates and to high packing density. 5. capillary water absorption allow to consider all mixtures as high quality, but the mixtures with slag addition promotes high reduction of water absorption compared to references, from 45 to 60%, followed by fly ash mixtures, also with reductions up to 45%. 6. in carbonation resistance, the cement dosage has an important positive influence on reducing carbonation depth, being reference mixtures those with less carbonation; however, mixtures with slag addition have similar or even better effect in reducing carbonation than references. 7. chloride diffusion coefficient of mixtures with additions proves a high influence of pozzolanic effect and the corresponding maturity of the matrices with age, since chloride migration reduce until 65% from 28 to 56 days of age; the additions of fly ash and of pozzolans are very advantageous for resistance to chloride migration, being the latter the best, since it reduces the chloride coefficient by 80 to 85%; contrarily, slag addition proves a different effect on tested mixtures, since it increases the chloride coefficient from 25 to 60% in comparison to references, and it does not improve with age. overall, the evolution of conventional cementitious mixtures to eco-efficient mixtures with less cement and addition does not decrease the performance of materials, on the contrary, depending on the additions, it may result in significant improvements of the mechanical, time dependent and durability performances of concrete mixtures. acknowledgements this work is co-financed by the european regional development fund (erdf), through the partnership agreement portugal2020 operational program for competitiveness and internationalization (compete2020), under the project poci-01-0247-feder033523 "postejo 4.0: inovação para a diversificação e a exportação". references [1] l. carvalho. experimental characterization of the deferred properties and durability of concrete with low cement dosage, master thesis, isec-polytechnic institute of coimbra, 2016. [2] t. proske, s. hainer, m. rezvani, et al. eco-friendly concretes with reduced water and cement content mix design principles and application in practice. construction and building materials 67:413-21, 2014. https: //doi.org/10.1016/j.conbuildmat.2013.12.066. 576 https://doi.org/10.1016/j.conbuildmat.2013.12.066 vol. 33/2022 mechanical and durability performance of cementitious matrices [3] s. fennis, j. walraven, j. uijl. the use of particle packing models to design ecological concrete, heron, 54(2):185-204, 2009. http://heronjournal.nl/54-23/5.pdf. [4] h. costa, h. alves, e. freitas, e., et al. mechanical performance of concrete with low cement dosage. in national meeting of structural concrete be2016, coimbra, 2016. [5] s. a. a. m. fennis, j. c. walraven, j. a. den uijl. compaction-interaction packing model: regarding the effect of fillers in concrete mixture design. materials and structures 46(3):463-78, 2012. https://doi.org/10.1617/s11527-012-9910-6. [6] s. fennis. design of ecological concrete by particle packing optimization. phd thesis, delft university of technology, 2011. [7] en 1015-3. methods of test for mortar for masonry, part 3: determination of consistence of fresh mortar (by flow table), 1999. [8] en 1015-7. methods of test for mortar for masonry, part 7: determination of air content of fresh mortar, 1998. [9] en 1015-11. methods of test for mortar for masonry, part 11: determination of flexural and compressive strength of hardened mortar, 1999. [10] en 1015-13. methods of test for mortar for masonry, part 13: determination of dimensional stability of hardened mortars, 1995. [11] e 397. hardened concrete. determination of the modulus of elasticity of concrete in compression (in portuguese). lnec, lisbon, 1993. [12] e393. concrete. determination of the absorption of water through capillarity (in portuguese). lnec, lisbon, 1993. [13] nt build 357. concrete, referring materials and protective coating: carbonation resistance, 1989. [14] nt build 492. concrete, mortar and cement-based repair materials: chloride migration coefficient from non-steady-state migration experiments, 1999. [15] browne, r. d. field investigations: site & laboratory tests: maintenance repair and rehabilitation of concrete structures, ceec, lisbon, 1991. 577 http://heronjournal.nl/54-23/5.pdf https://doi.org/10.1617/s11527-012-9910-6 https://doi.org/10.14311/app.2022.33.0070 acta polytechnica ctu proceedings 33:70–77, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague natural fiber-reinforced light-weight cement blocks prepared from waste for sustainable development latda chandenga, thanawat meesaka, chakrapan tuaktab, tidarut wisuthseriwongb, tamon uedac, oratai jongprateepa, ∗ a kasetsart university, faculty of engineering, department of materials engineering, 50 ngamwongwan road, lat yao, chatuchak, 10900 bangkok, thailand b kasetsart university, faculty of engineering, department of civil engineering, 50 ngamwongwan road, lat yao, chatuchak, 10900 bangkok, thailand c shenzhen university, guangdong provincial key laboratory of durability for marine civil engineering, college of civil and transportation engineering, shenzhen, guangdong, 518054, china ∗ corresponding author: fengotj@ku.ac.th abstract. low density of the light-weight cement blocks offers an advantage in terms of dead load reduction, which is advantageous in structural design-reduction of size and numbers of loadcarrying structural components is possible. production of cement blocks, however, generally requires ordinary portland cement (opc) which creates harmful environmental impacts. utilization of waste as alternative raw materials for cement production is a route to alleviate the problem. this research aims at synthesizing eco-friendly cement-like material for production of light-weight cement blocks. the cement-like material were prepared from eggshells, cockleshells, and rice husk ash (rha). with high content of calcium oxide and silica, eggshells and cockleshells are potential sources of calcium, whereas rha is a good source of silicon. additionally, a fuel-efficient solution combustion technique was employed in synthesis of the cement-like material. phase identification analysis of the synthesized powder indicated that tri-calcium silicate (c3s), di-calcium silicate (c2s), tri-calcium aluminate (c3a), and tetra-calcium alumino ferrite (c4af), which are main constituents of opc, were obtained. to fabricate eco-friendly light-weight cement blocks, the synthesized cement-like material were mixed with cement, water, and additional rha and cast into blocks. the optimal compressive strength and density of the cement blocks were in comparable range with the standard light-weight concrete defined by thai industrial standards institute (tisi) and american concrete institute (aci 213,2001). with jute fiber reinforcement, enhanced compressive strength of 20% was achieved, while elimination of spalling after compression test was clearly evident, implying a more ductile failure. keywords: fiber reinforcement, light-weight cement block, waste. 1. introduction as a predominantly agricultural country, agricultural sector in thailand plays a significant role in economic contribution. with increasing world population and growing demand for agricultural products, considerable amounts of agricultural residues are produced annually. it has been reported that approximately 5 million tons of rice husk are produced in thailand annually [1]. residues from poultry industry, one of thailand’s promising agri-food segments, is notable. production of eggshells in thailand is estimated to be around 650,000 tons per year [2]. in addition to rice husk and eggs, residue produced from aquaculture is abundant, with approximate amount of 10 thousand tons per year for cockleshells [3]. these remainders are often transformed into waste, disposed in landfill sites and possibly create pollution [4–7]. efficient utilization of the aforementioned waste is vital to both the economy and environment. it is well recognized that construction industry is the major consumer of natural resources. to preserve natural resources, emerging technology, such as innovative utilization of waste for construction, is commendable. since key constituents of cement comprise calcium silicate and calcium aluminate, wastes containing extensive amounts of calcium and silicon have been considered. according to numerous studies, rice husk ash (rha) contains significant silica content, with the amount exceeding 90 wt% [8, 9]. cockleshells and eggshells are reported to contain high level of calcium carbonate, in the range between 95 and 99% [10–13]. hence, rice husk ash, eggshells, and cockleshells can be potential sources of silicon and calcium used as cement-like material. along with waste utilization, an efficient route of waste management and reduction of natural resource consumption, production of eco-friendly cement-like material by fuel-efficient techniques can also be a route to save the environment. reduction of carbon dioxide emissions can provide a positive impact 70 https://doi.org/10.14311/app.2022.33.0070 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 waste light-weight cement blocks components components ratio tri-calcium silicate (c3s) and di-calcium silicate (c2s) 0.75 tri-calcium aluminate (c3a) 0.12 tetra-calcium alumino ferrite (c4af) 0.10 gypsum 0.03 table 1. the components of cement-like material. to the environment. conventional production of cement requires heating process of raw materials, also called clinker, at the temperature ranging between 1400 and 1600◦c. with high temperature calcination, enormous fossil fuel consumption and massive carbon dioxide emission are intensified. it has been reported that 0.81 kg of cement production produces co2 1 kg on average [14]. carbon dioxide emission from cement production is a result of processing steps, including 7% from quarrying, 5% from clinker grinding, 3% transportation. calcination of the clinker is accounted for 85% of carbon dioxide emission [15]. the natural resources and environmental policy and planning of thailand reported that cement industry is responsible for carbon dioxide emission of more than 12 million tons per year [16]. solution combustion technique is a powder synthesis technique capable of producing fine ceramic powders with controlled chemical composition. the technique involves self-propagating exothermic reaction. in addition to high-quality powder production, this method offers advantages in terms of low fossil fuel consumption and minimal processing time [17]. since combustion reaction is initiated at the temperature lower than 400◦c, and low calcination temperature of 900◦c is required for production of cement-like material, considerably lower fuel consumption and diminished carbon dioxide emission is substantially evident. for practical applications, lightweight cement blocks are required to maintain compressive strength following industrial standards. thai industrial standards institute (tisi) type c12 demands minimum compressive strength of 2.5 mpa [18]. in addition to adequate strength, high durability and good integrity of the cement blocks are desired. poor installation of construction component, excessive stress, and environmental factors often cause disintegration of the materials. flaking and pitting are commonly observed. with addition of fiber reinforcement, alleviation of spalling is possible. fiber reinforcement can generally enhance mechanical properties of materials through crack abridgement. it has been reported that jute fiber is an effective reinforcing material to improve ductility, compressive strength, flexural strength, and tensile strengths in concrete [19, 20]. this research aims at utilizing waste in synthesis of eco-friendly cement-like materials. properties of fiber-reinforced light-weight cement blocks prepared from the cement-like materials were also examined. 2. experimental procedures 2.1. preparation of cement-like material an initial step to synthesis of cement-like material involves preparation of wastes. to obtain fine particle, with sizes smaller than 45 micrometers, rice husk ash (dhebkaset industrial co., ltd) was ground by ball milling for 2 hours and screened through a 325mesh sieve. eggshells and cockleshells were collected from local markets in bangkok, calcined at 900◦c for 3 hours and ball milled. solution combustion technique was employed in synthesis cement-like material, consisting of dicalcium silicate (c2s), tri-calcium silicate (c3s), tricalcium aluminate (c3a), and tetra-calcium alumino ferrite (c4af). di-calcium silicate (c2s) and tricalcium silicate (c3s) were produced by dissolving cockleshells or eggshells and rice husk ash (rha) in nitric acid 70% (hno3, rci labscan) with the weight ratio of cockleshell:rha equal to 0.84 : 0.41. for preparation of c2s and c3s using eggshells, the weight ratio of eggshell:rha of 0.85 : 0.41 was used. glycine (c2h5no2, daejung) and deionized water was added to the mixture to obtain aqueous solution with concentration close to 0.2 m, which was subsequently heated at 400◦c. combustion reaction was initiated and resulted in as-synthesized powders. the powder was collected and calcined at 900◦c for 3 hours. c3a and c4af powders were prepared in a similar approach. for c3a, cockleshells or eggshells were mixed with aluminium nitrate nonahydrate (al(no3)3, daejung) at weight ratio close to 0.73 : 2.77. c4af was attained by preparing mixture of cockleshells or eggshells, aluminium nitrate nonahydrate, and 95% ferric oxide red (fe2o3, lobachemie) at ratio of 0.81 : 2.36 : 0.51 by weight. to obtain chemical compositions similar to those of the opc, the synthesized powder of tricalcium silicate(c3s) and di-calcium silicate(c3s), tri-calcium aluminate (c3a), tetra-calcium alumino ferrite (c3af) and gypsum were mixed together with weight ratio of cs : c3a : c4af : gypsum equal to 0.75 : 0.12 : 0.1 : 0.03, as shown in table 1. 2.2. casting of eco-friendly lightweight cement blocks in order to utilize high content of waste and to minimize usage of opc, lightweight cement blocks con71 l. chandeng, t. meesak, ch. tuakta et al. acta polytechnica ctu proceedings components of light-weight cement blocks content of fiber cement-like opc rha water to binder ratio[%w/v] [wt%] [wt%] [wt%] 1 20 20 60 0.8, 0.9, 1, 1.1, 1.2 1 25 25 50 0.8, 0.9, 1, 1.1, 1.2 1.5 20 20 60 0.8, 0.9, 1, 1.1, 1.2 1.5 25 25 50 0.8, 0.9, 1, 1.1, 1.2 2 20 20 60 0.8, 0.9, 1, 1.1, 1.2 2 25 25 50 0.8, 0.9, 1, 1.1, 1.2 table 2. components of light-weight cement blocks consisting of cement-like material, opc, rice husk ash and natural jute fiber with water to binder ratio of 0.8 to 1.2 and autoclaving 6 hours after curing in water 7 days. figure 1. a) x-ray diffraction pattern of rice husk ash, b) x-ray diffraction pattern of cockleshells, c) x-ray diffraction pattern of eggshells taining 50 − 60 wt% of rha were fabricated. the mixtures containing cement-like material, opc, rha at weight ratio of cement-like material:opc:rha equal to 25 : 25 : 50 and 20 : 20 : 60, respectively, were prepared. to examine optimal water content, the mixtures were mixed with water at the water to binder ratio (w/b) of 0.8, 0.9, 1, 1.1 and 1.2. to examine effects of fiber reinforcement, 1 cm-long jute fibers were added to the mixtures at the content of 1, 1.5 and 2% w/v. the mixtures were casted in a 2.5 × 2.5 × 2.5 cm acrylic mold, cured in water for 7 days and autoclaved for 6 hours. components of the fiber-reinforced cement blocks are shown in table 2. 2.3. characterization the examination of chemical composition and phase identification of the raw waste materials, such as rice husk ash, cockleshells, and eggshells, as well as the cement-like material, were conducted by x-ray diffraction (philips, xpert). compressive strength of the lightweight cement blocks was determined by using a universal testing machine (hounsfield, h50ks), whereas optical micrographs were obtained from the optical microscope (lica, dm2700m). 3. results and discussion 3.1. phase identification of raw waste materials x-ray diffraction was employed in phase identification of rice husk ash, cockleshells, eggshells. according to x-ray diffraction patterns shown in figure 1, rice husk ash contained a single phase of silica (sio2, jcpds 01-074-9378). observations from this study was consistent with numerous researches [21–23]. the x-ray diffraction patterns of cockleshells and eggshells were similar in nature. the patterns, as shown in figure 1, contained prominent peaks corresponding to calcium hydroxide (ca(oh)2, jcpds 01-084-1263) and calcium carbonate (caco3, jcpds 01-085-1108). with high silica content and high calcium content, rice husk ash, cockleshells and eggshells had great potential for usage as raw materials for producing cement constituents such as tri-calcium silicate (c3s) and di-calcium silicate(c2s), tri-calcium aluminate (c3a) and tetra-calcium alumino ferrite (c4af). 3.2. phase identification of cement and cement-like material as mentioned in the previous section, high silica content and high calcium content facilitate the use of rice husk ash and cockleshells in the production of c3s, c2s, c3a, and c4af. in this section, xray diffraction was also employed in phase identification of the synthesized powders. x-ray diffraction pattern of the synthesized powders revealed prominent peaks corresponding to tri-calcium silicate (ca3sio5, jcpds 00-031-0301), di-calcium silicate (ca2sio4, jcpds 00-029-0369), tri-calcium aluminate (ca3al2o6, jcpds 00-008-0005) and tetracalcium alumino ferrite (ca4al2fe2o10, jcpds 00011-0124). the pattern, as showed in figure 2, was also similar to the diffraction pattern obtained from 72 vol. 33/2022 waste light-weight cement blocks figure 2. x-ray diffraction pattern of the ordinary portland cement (opc) and cement-like material containing tri-calcium silicate (ca3sio5), di-calcium silicate (ca2sio4), tri-calcium aluminate (ca3al2o6) and tetra-calcium alumino ferrite (ca4al2fe2o10) figure 3. a relationship between bulk density and w/b ratio of jute fiber-reinforced cement blocks containing rha50% and rha60%. opc. the results suggested that the cement-like material synthesized by solution combustion technique might be alternative materials used for substitution of ordinary cement. 3.3. effects of water content on density of lightweight cement blocks water-to-binder ratio (w/b) is one of the most critical parameters influencing properties of construction components containing cement, specifically workability, strength, and durability. determination of the w/b is therefore crucial for quality control during lightweight cement blocks production and general quality assurance purposes [24, 25]. in this study, the mixture containing synthesized cement-like material, opc, rice husk ash, and jute fibers were mixed with water at the w/b ratios ranging from 0.8 to 1.1, and cast into cement blocks. after 7 days of curing in water and autoclave, the cement blocks were tested for their density. the results, as shown in figure 3, revealed density values ranging from 0.873 to 1.050 g/cm3 and from 0.796 to 1.013 g/cm3 in the fiber-reinforced cement blocks containing 50 and 60 wt% rice husk ash, respectively. high water content diminished bulk density. this is attributed to formation of capillary water from excessive water content. the water subsequently evaporates and leads to formation of capillary pores. it was also found that the cement blocks containing high rha yielded lower density. this is attributed to low bulk density of rice husk ash (dhe73 l. chandeng, t. meesak, ch. tuakta et al. acta polytechnica ctu proceedings figure 4. a relationship between water to binder ratio and compressive strength of jute fiber-reinforced cement blocks containing rha50% and rha60%. figure 5. a relationship between compressive strength and jute fiber contents of cement blocks containing rha50% and rha60%. bkaset industrial co., ltd) of 0.33 g/cm3. according to type c12-thai industrial standards institute (tisi 2601-2556), it is required that the lightweight concrete must have density lower than 1.2 g/cm3. the overall density values of the cement blocks obtained in this study were in an acceptable range according to tisi 2601-2556. 3.4. effects of water content on compressive strength of lightweight cement blocks the results from compressive strength measurements of the fiber-reinforced cement blocks indicated that the strength values ranging from 2.6 to 5.1 mpa and 2.1 to 4.1 mpa was achieved in the cement blocks containing 50 and 60 wt% rice husk ash, respectively. in this study, the optimal w/b ratio for enhancement of compressive strength was 0.8. with high w/b ratio, compressive strength was diminished. minimal compressive strength was associated with low density, which was the result of excessively high water content. in addition, the results, as shown in figure 4, also revealed that higher content of rice husk ash could result in reduction of compressive strength. with high rha, content of cement and cement-like material, which play a role as binding materials and reagent for hydration reactions, was minimized. uneven and discontinuous distribution of the cementitious materials could occur, which possibly cause low compressive strength. 74 vol. 33/2022 waste light-weight cement blocks figure 6. lightweight cement blocks after compression: a) rha 50% w/b 0.8 autoclaved 6 hours (no jute fiber), b) rha 50% w/b 0.8 autoclaved 6 hours (with 1.5% jute fibers). figure 7. optical microscographs image (x5) of lightweight cement blocks after compression test. 3.5. effects of fiber contents on compressive strength of lightweight cement blocks in this study, compressive strength of cement blocks with w/b lower than 1.1 were higher than the minimum compressive strength values required by thai industrial standards institute (tisi 2601-2556)-type c12. with acceptable range of both density and compressive strength, jute fiber-reinforced cement blocks tended to have great potential for practical utilization. in the previous sections, it was found that the optimal water to binder ratio to achieve acceptable density while retaining high compressive strength was 0.8. in this section, optimal content of jute fibers for enhancement of compressive strength was examined. for the cement blocks containing 50 wt% rha, the minimal and optimal compressive strength was observed in the un-reinforced cement blocks and the reinforced cement blocks with 1.5% jute fiber, respectively. a similar trend was observed in the cement blocks containing 60 wt% rha, as shown in figure 5. the results indicated that with an optimal content of fiber reinforcement, compressive strength could increase from 4 to 5.12 mpa and from 3.5 to 4.13 mpa in cement blocks containing 50 and 60 wt% rha, respectively. an increment of 18 to 28% of compressive strength was consistent with numerous studies [26]. fibers with high tenacity, such as jute fibers, can be substantially beneficial for improvement of mechanical properties of cement blocks. in addition, high homogeneity mixing of short random fibers can increase the integral adhesion force within the cement blocks, resulting in reduction of spalling in cement blocks when subjected to compression, as showed in figure 6. observation from this study was concurrent with s. p. kundu et al, which revealed that the un-reinforced samples were disintegrated after the maximum load was applied. on the contrary, for the samples with jute fiber reinforcement-cracks appeared without spalling even when the samples were subjected to the maximum load [27–29]. it has been generally accepted that mechanical property enhancement of fiber-reinforced composites can be attributed to crack bridging mechanism. as shown in figure 7, jute fibers functioned as bridges, joining cement matrix, which consequently resulted in good integrity of the samples subjected to compressive stress. despite the aforementioned advantages, excessively high fiber content did not result in improved compressive strength. according to m. zakaria et al, addition of fiber of more than 2% could lead to fiber ball-up and intermeshing while mixing, and resulted in inferior mechanical strength [19, 30, 31]. 4. conclusion eco-friendly fiber-reinforced light-weight cement blocks with acceptable density and compressive strength were successfully fabricated. cement-like material, an alternative material for opc which was used as binder for cement block production, was synthesized by solution combustion technique. in addition to utilization of waste, specifically cockleshells, eggshells and rice husk ash, the synthesis technique required low energy consumption, which could be integral part toward sustainable development. additionally, effects of processing parameters, including water to binder ratio, rice husk ash content and jute fiber content, were examined. to achieve 75 l. chandeng, t. meesak, ch. tuakta et al. acta polytechnica ctu proceedings the optimal compressive strength, water to binder ratio of 0.8 and 50% of rice husk ash were required in this study. by adding jute fiber with the amount of 1.5%, compressive strength enhancement as high as 28% could be achieved. acknowledgements all authors would like to extend sincere gratitude to aun/seed-net for financial support. equipment support from the department of materials engineering, faculty of engineering, kasetsart university, and raw materials from dhebkaset industrial co., ltd are gratefully acknowledged. references [1] s. nattawut, t. rinnaphat, c. raungrut, 2019. http: //raungrut.tunjai.com/index.php?uid=1&pid=23&me nutype=rboard&qid=60&fbclid=iwar3mxglgyks09cqy 2hx5vkksiilgqv8vzexcvkedyszo5nucakf80ex7fbk. 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[31] l. yan, n. chouw. natural frp tube confined fibre reinforced concrete under pure axial compression: a comparison with glass/carbon frp. thin-walled structures 82:159-69, 2014. https://doi.org/10.1016/j.tws.2014.04.013. 77 https://doi.org/10.1016/j.cemconres.2009.06.013 https://doi.org/10.1016/j.conbuildmat.2019.04.036 https://doi.org/10.1016/j.conbuildmat.2018.10.045 https://doi.org/10.1016/j.jobe.2019.100786 https://doi.org/10.1016/j.conbuildmat.2018.09.048 https://doi.org/10.1016/j.tws.2014.04.013 acta polytechnica ctu proceedings doi:10.14311/app.2019.25.0017 acta polytechnica ctu proceedings 25:17–20, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app compressive properties of auxetic structures with controlled stiffness of strut joints tomáš doktor∗, tomáš fíla, petr koudelka, daniel kytýř, ondřej jiroušek czech technical university in prague, faculty of transportation sciences, konviktská 20, 110 00 prague, czech republic ∗ corresponding author: doktor@fd.cvut.cz abstract. presented paper deals with experimental study on compressive properties of auxetics with controlled stiffness of strut joints. the variable strut joints properties were simulated by adding extra amount of material in the struts’ intersection regions. four groups of inverted honeycomb structures were prepared by multi-jet 3d printing and tested in quasi-static compression. the structure collapsed gradually, however after the first collapse, failure in entire cross-section occurred due to the brittle nature of the base material. the behavior up to the first collapse was consistent among the specimens within each group, while differed slightly subsequently. with higher reinforcement in the joints, results showed increasing stress at the first collapse (ultimate compressive stress) while the strain at the first collapse remained unchanged. the auxetic behaviour became less significant with increasing joints’ reinforcement. keywords: auxetics, cellular materials, quasi-static testing, optimisation, additive manufacturing. 1. introduction it has been already shown that cellular solids are able to absorb tremendous amounts of deformation energy [1, 2] during impacts with the possibility to introduce strain-rate dependent characteristics into their deformation response [3]. however, certain applications, such as blast and flying debris protection systems, may require material with relatively high strength in addition to excellent deformation absorption capabilities. such complex mechanical characteristics can be achieved without alteration of the existing lightweight alloys (such as aluminium based materials) used for the material’s production when the microstructure is constructed in such a way to exhibit negative poisson’s ratio [4]. mechanical characteristics of such auxetic lattices are given not only by overall geometrical arrangement of struts in the unit cell and their connectivity but also by deformation properties of the strut joints. there are two main mechanisms of the lattice behaviour during mechanical deformation. based on the struts’ connectivity the deformation is stretching dominated for lattices which comply with truss assumptions or bending dominated in lattices with incomplete connectivity [5]. hence the deformation and energy absorption properties are highly influenced by the bending in the struts’ connections. influence of all these factors on effective mechanical properties in both elastic and plastic regime has to be thoroughly evaluated and taken into account in analytical and numerical optimisation. in this paper a parametric study on influence of the strut joint stiffness on the effective deformation response of various auxetic lattices is presented. 2. materials and methods 2.1. specimen design and manufacturing auxetic lattices exhibiting negative poisson’s ratio were designed in parametric modeller solidworks (dassault systémes corp., france). three-dimensional inverted honeycomb was selected which exhibits auxetic behaviour in all three dimensions. design of the tested auxetic structure was based on our previous study [6], in which first preliminary results were presented. only entirely auxetic structures were used in contrary to the planar auxetics. the porosity and ratio between cell size and strut cross-section (square shaped) was accommodated for a better stiffness and strength. at the top and bottom plane plate was added in to ensure a better force transfer between the loading grips and the specimens. design parameters of the specimens are listed in table 1, geometry of the unit cell is depicted in fig. 1. specimen height [mm] 26 specimen cross-section [mm] 16.5 × 16.5 number of unit cells [−] 3 × 3 × 3 strut cross-section [mm] 0.75 × 0.75 table 1. design parameters common parameters. different joint stiffness was achieved by differently sized spheres placed in the intersections of the struts. three different radii were selected. upper limit value of the spheres’ radius (1.875 mm) was selected to avoid contact between neighboring spheres. the lower limit (0.938 mm) was selected with respect to a possible observation of the reinforcing effect, as smaller spheres 17 http://dx.doi.org/10.14311/app.2019.25.0017 http://ojs.cvut.cz/ojs/index.php/app t. doktor, t. fíla, p. koudelka et al. acta polytechnica ctu proceedings figure 1. geometry of the unit cell. would be entirely hidden in the intersection of the prismatic struts. parameters of the joints’ reinforcement are listed in table 2. group sphere radius [mm] porosity [%] 0 78.8 1 0.75 76.9 2 0.9375 72.4 3 1.125 64.9 table 2. design parameters joint reinforcement variations. stl models which were used for the 3d printing of tested samples are depicted in figure 2. the specimens were manufactured by direct 3d printing from uvcurable acrylic material visijet ex200 (3d systems, usa) using 3d printer pro jet hd3000 (3d systems, usa) in high definition mode. resolution in high definition mode was 387 × 387 × 300 dpi. figure 2. stl models for 3d printing. 2.2. compression tests for the compression experiments a custom loading device was used equipped with force transducer with loading capacity 2 kn (u9b, hottinger baldwin messtechnik gmbh, germany). the loading was provided by a stepper motor controlled by linuxcnc software solution used with real-time gnu/linux operating system [7]. the tests were displacement driven with loading rate 20µm · s−1. contactless displacement measurement was employed based on digital image correlation (dic) [8, 9]. the loading scene was captured by a ccd camera (manta g504b, alied vision technologies gmbh, germany) attached to a figure 3. loading scene with tracking features (left) and tracked displacement paths (right). black background was inverted due to a better visibility. bi-telecentric zoom lens (tzcr 072, optoengineering, italy). the image data was captured with a custom software tool based on opencv library [10]. zoom ratio 0.25 × was used for all compressive tests, which provided pixel-size 13.7 µm. a set of tracking features was selected in the first (undeformed) loading scene which were arranged into 7 columns and 15 rows to ensure placement of the features in solid material of the struts. the array of loading features is depicted in figure 3 (left). the loading rate together with the frame rate 2 fps ensured a sufficient number of points in the obtained loading curve. figure 3 (right) shows displacement paths tracked by dic in selected specimen from group 2. 2.3. evaluation the measured displacements and forces were used to calculate engineering strain �eng and stress σeng which were converted into true strain �true and true stress σtrue using formulae (1) and (2), respectively. �true = ln(1 + �eng ) (1) σtrue = σeng · (1 + �eng ) (2) based on measured strain in transversal and longitudinal direction, poisson’s ratio was calculated according to formula (3). µ = − �x �y (3) to exclude values which correspond to both initial settlement and post-yield region after the first plastic collapse, the function of poisson’s ratio was limited only inside these limit values of strain. moreover, to assess the influence of joints’ stiffness on the mechanical behaviour, ultimate compressive stress of the tested structures was evaluated as well as strain corresponding to the first plastic collapse. 18 vol. 25/2019 compressive properties of auxetics with controlled joint stiffness group σm [mpa] �y1 [1] µmean [1] 0 0.876 ± 0.124 0.0383 ± 0.0032 −0.180 ± 0.040 1 0.887 ± 0.134 0.0348 ± 0.0029 −0.219 ± 0.058 2 0.942 ± 0.138 0.0346 ± 0.0030 −0.142 ± 0.030 3 1.168 ± 0.054 0.0352 ± 0.0027 −0.047 ± 0.065 table 3. average results in tested groups. stress and strain at the first collapse. average poisson’s number. 0 0.2 0.4 0.6 0.8 1 1.2 1.4 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 t ru e s tr e s s [ m p a ] logarithmic strain [-] 0.1 0.2 0.3 0.4 0.5 figure 4. stress-strain curves of specimens in group 0. 0 0.2 0.4 0.6 0.8 1 1.2 1.4 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 t ru e s tr e s s [ m p a ] logarithmic strain [-] 1.1 1.2 1.3 1.4 figure 5. stress-strain curves of specimens in group 1. 3. results 3.1. stress-strain curves a series of uniaxial compressive tests was performed on custom designed three-dimensional inverted honeycomb structures. from the captured loading scenes the evolution of the deformation of the tested structures is assessed. unfortunately, the base material exhibited a high brittleness which caused failure after the yield strain of the structure. after this failure the loading continued with similar stress-strain behaviour but with a slightly lower stresses at the consequent collapses. true stress vs. logarithmic strain diagrams are depicted in figures 4 to 7 for groups 0 to 3 respectively. to compare the main mechanical properties of the tested structures and to assess influence of the different reinforcement of the strut joints, mean values among the groups were evaluated with their standard 0 0.2 0.4 0.6 0.8 1 1.2 1.4 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 t ru e s tr e s s [ m p a ] logarithmic strain [-] 2.1 2.2 2.3 2.4 2.5 figure 6. stress-strain curves of specimens in group 2. 0 0.2 0.4 0.6 0.8 1 1.2 1.4 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 t ru e s tr e s s [ m p a ] logarithmic strain [-] 3.1 3.2 3.3 3.4 3.5 figure 7. stress-strain curves of specimens in group 3. deviations. in table 3 are listed mean ultimate compressive strength (stress at the first collapse), strains corresponding to the first collapse in each group of samples and mean poisson’s ratio in the region up to the first collapse. 4. conclusions an experimental study on deformation response of 3d printed auxetic structures was carried out to evaluate influence of joint properties which were physically simulated by adding an extra amount of material in the struts’ connections. the loading setup was equipped with a custom contactless strain measurement system which enabled a precise evaluation of the stress-strain curve. the obtained curves exhibited consistent behaviour among each group (i.e. sphere radius) in terms of stiffness in the linear part, ultimate compressive strength and strain at the first collapse. 19 t. doktor, t. fíla, p. koudelka et al. acta polytechnica ctu proceedings the stress level at the first collapse increases with increasing stiffness in the joints, while the strain at the first collapse remains unchanged. the auxetic nature exhibits slight increase when the spheres with the lowest radius are added in the joints’ interconnections. then, with increasing spheres’ radius the auxetic behaviour is less significant. these results show promising possibility of optimisation of the cellular structures based on controlled stiffness in the strut joints. in the further results this knowledge will be employed for optimisation of the auxetics using different materials in the interconnection regions. acknowledgements the financial support of the czech science foundation (project no. 19-23675s). references [1] q. chen, n. m. pungo. in-plane elastic buckling of hierarchical honeycomb materials. european journal of mechanics a/solids 34:120–129, 2012. doi:10.1016/j.euromechsol.2011.12.003. [2] r. harb, e. taciroglu, n. ghoniem. partitioning of elastic energy in open-cell foams under finite deformations. acta materialia 61(5):1454–1468, 2013. doi:10.1016/j.actamat.2012.11.022. [3] k. a. dannemann, j. lankford jr. high strain rate compression of closed-cell aluminium foams. materials science and engineering 293(1):157–164, 2000. doi:10.1016/s0921-5093(00)01219-3. [4] r. lakes. foam structures with a negative poisson’s ratio. science 235(4797):1038–1040, 1987. doi:10.1126/science.235.4792.1038. [5] v. s. deshpande, m. f. ashby, n. a. fleck. foam topology: bending versus stretching dominated architectures. acta materialia 49(6):1035–1040, 2001. doi:10.1016/s1359-6454(00)00379-7. [6] p. koudelka, o. jirousek, t. fila, t. doktor. compressive properties of auxetic structures produced by direct 3d printing. materiali in tehnologije 50(3):311–317, 2016. [7] v. rada, t. fila, p. zlamal, et al. multi-channel control system for in-situ laboratory loading devices. acta polytechnica proceedings 18:15–19, 2018. doi:10.14311/app.2018.18.0015. [8] b. d. lucas, t. kanade. an iterative image registration technique with an application to stereo vision. in: proceedings of imaging understanding workshop pp. 121–130, 1981. [9] i. jandejsek, j. valach, d. vavrik. optimization and calibration of digital image correlation method. in: proceedings of experimental stress analysis 2010 pp. 121–126, 2010. [10] g. bradski. the opencv library. dr dobb’s journal of software tools 2000. http://www.drdobbs.com/ open-source/the-opencv-library/184404319. 20 https://doi.org/10.1016/j.euromechsol.2011.12.003 https://doi.org/10.1016/j.actamat.2012.11.022 https://doi.org/10.1016/s0921-5093(00)01219-3 https://doi.org/10.1126/science.235.4792.1038 https://doi.org/10.1016/s1359-6454(00)00379-7 https://doi.org/10.14311/app.2018.18.0015 http://www.drdobbs.com/open-source/the-opencv-library/184404319 http://www.drdobbs.com/open-source/the-opencv-library/184404319 acta polytechnica ctu proceedings 25:17–20, 2019 1 introduction 2 materials and methods 2.1 specimen design and manufacturing 2.2 compression tests 2.3 evaluation 3 results 3.1 stress-strain curves 4 conclusions acknowledgements references doi:10.14311/app.2021.30.0024 acta polytechnica ctu proceedings 30:24–29, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app modeling time-dependent deformations of concrete members subjected to symmetric and asymmetric drying pavel horák∗, petr havlásek czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: pavel.horak@fsv.cvut.cz abstract. the time-dependent behavior of concrete members with uniform cross-section subjected to symmetric drying can be predicted with one of many models anchored in the design codes. since these models use the cross-sectional approach, only the average quantities can be obtained. more complex phenomena can be investigated by means of a coupled hygro-mechanical finite element analysis. owing to insufficient or missing experimental data, the input parameters for the moisture transport model are often tuned to match the development of axial shrinkage or moisture loss measured on small-scale laboratory specimens. in this paper it is assessed whether this approach leads to a unique set of parameters of the bažant-najjar model for moisture diffusion. additionally, for the selected sets of parameters the analysis is repeated with modified size and/or boundary conditions leading to asymmetric drying and the results are discussed. keywords: concrete, drying shrinkage, finite elements, moisture transport, parameter identification. 1. introduction and motivation creep and shrinkage are complex time-dependent phenomena in concrete whose magnitude and evolution is influenced by many factors. the distinction between creep and shrinkage, which are both related to delayed strains, is that creep is caused by loading. at constant temperature the creep deformation can be additively decomposed into basic creep and drying creep. basic creep is the deformation of hygrally sealed concrete while drying creep is an additional deformation caused by drying. in a similar fashion, autogenous and drying shrinkage refer to the volume change (typically contraction) of sealed and drying concrete, respectively. in normal-strength structural concretes with w/c ≥ 0.4, the autogenous shrinkage is typically negligible (≈ 40 × 10−6) in comparison to the magnitude of drying shrinkage (≈ 780 × 10−6 at 40% relative humidity according to aci 209). this time-dependent behavior of typical concrete members can be estimated using the prediction models (e.g. b3, b4, fib model code) or the formulae from the design codes (e.g. aci 209, eurocode 2). however, since these models operate on the cross-sectional level, only the average quantities can be obtained. in order to describe more complex phenomena, e.g. tensile cracking produced by stress gradients caused by nonuniform drying or the structural response to non-symmetric drying, a more versatile approach is a necessity. one of the options is a coupled hygro-mechanical finite element analysis. the bottleneck of this approach resides in often insufficient number of experiments which are necessary for proper calibration of the complex constitutive models. to overcome this obstacle, certain parameters are set to their recommended values or the parameters are adjusted to reproduce the observed macroscopic behavior. to illustrate this, in the recent paper [1], this approach was used by several research groups to calibrate the transport model based on the axial shrinkage and moisture loss of a standard laboratory specimen without any information on the distribution of relative humidity. the primary objective of this paper is to assess whether such approach yields a unique set of material parameters and what macroscopic measurements can narrow the range of the identified values. 2. methodology all results presented further in this paper were obtained using a one-way coupled hygro-mechanical simulations in open-source finite element package oofem [2]. the constitutive models described in more detail in the following sections were calibrated on creep and shrinkage data reported by bryant [3]. the parameters of the structural material model are then fixed while the parameters in the model for moisture transport are subject to optimization under various assumptions. the identified parameters are consequently utilized for predicting the curvature of an infinite slab subjected to asymmetric drying. 2.1. material models concrete drying is in the present study described by a widely recognized model proposed by bažant and najjar [4]. the governing equation for the diffusion of water vapor was derived under the assumption of 24 http://dx.doi.org/10.14311/app.2021.30.0024 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 modeling of concrete subjected to symmetric and asymmetric drying constant moisture capacity k (slope of the desorption isotherm). when the moisture sink is neglected the equation reads ∂h ∂t = ∇ · (c(h)∇h) (1) where ∇h is the gradient of relative humidity and c(h) is the humidity-dependent diffusivity. this dependence is for the cementitious materials highly nonlinear and can be approximated as c(h) = c1 ! "α0 + 1 − α0 1 + # 1−h 1−hc $n % & (2) where c1 is the maximum diffusivity at h = 1, α0 determines the ratio between minimum diffusivity at h = 0 and c1, and parameters hc and n describe the relative humidity and the rate of the transition. this model was embedded into the fib model code 2010 [5] which recommends the following values α0 = 0.05, hc = 0.8, n = 15, while c1 is to be determined from the compressive strength. in the simulations, the initial value of relative humidity is set to h = 0.98 which is the typical value for the sealed conditions and normal-strength concrete with negligible self-desiccation. the ambient relative humidity is idealized as constant and is prescribed using the mixed boundary condition which uses the surface factor f to relates the humidity flux jh with the difference between the relative humidity at the boundary and the environment jh = f · (henv − h) (3) in each time step the moisture transport is followed by the structural sub-problem which utilizes the computed field of relative humidity. the constitutive vicoelastic model is the modified [6] microprestresssolidification (mps) theory [7]. under constant temperature and hygrally sealed conditions the material model reduces to the b3 model [8]. changes in relative humidity give rise to additional compliance, referred to as drying creep or the pickett effect, and to volume changes–shrinkage or swelling. the model has been modified to minimize the size-effect on drying creep which is controlled by parameter k3. shrinkage strain and relative humidity are linearly linked via their rates, ε̇sh = kshḣ (4) where ksh is a material parameter usually treated as humidityand age-independent constant. the model can be enhanced by tensile softening but this feature is not activated in the present study. 2.2. modeling strategy despite the unidirectional nature of the transport problem, the analysis was performed in two dimensions to provide the subsequent structural analysis with the distribution of relative humidity. the mechanical problem was carried out under the assumption of plane stress which allowed to run the computations very efficiently and repeatedly as described later. the assumption of plane stress is very realistic in the case of the experimental data. the parametric study is done on an idealized plane section cut out from an “infinite slab” drying from one or two surfaces. yet, the model which is hereafter referred to as “slab”, neglects the stresses perpendicular to the section and its behavior corresponds to a beam with prismatic cross-section, sealed lateral sides and drying from top and/or bottom surfaces. the effect of self weight is omitted in the study. the finite element meshes of the two sub-problems were identical and were composed of 60 quadrilateral elements with linear approximation of humidity and displacement. the meshes were refined near the drying surfaces. all problems were solved in 284 time steps with increasing length. as mentioned earlier, the material parameters were calibrated on experimental data by bryant and vahanavikkit [3]. their study was focused on creep and shrinkage of structural concrete (fc = 50.1 mpa, w/c = 0.47) measured on prismatic members with cross-section ranging from 100×100 to 400×400 mm. the specimens were either entirely sealed to investigate basic creep or partially sealed to study drying creep and shrinkage. drying at 60% relative humidity began at the age of 8 days. the majority of the creep tests imposing 7 mpa compression started at the age of 14 days. additional tests were done to identify aging. all loaded specimens had a companion without external loading. the constant room temperature was maintained throughout the experiment. calibration began with the basic creep compliance function. the parameters q1 − q4 of the b3 model were adopted from [9] and their values were q1 = 9, q2 = 75, q3 = 28 and q4 = 6.5 all in 10−6/mpa. drying creep and shrinkage was calibrated on experimental data of prism 150 × 150 mm drying from two lateral sides (i.e. 150 mm thick infinite wall). this specimen is the largest one which has reached the ultimate values and for this reason was selected as reference. with crudely calibrated parameters of the transport model the shrinkage constant was identified as ksh = 1.9 × 10−3, and drying creep parameter k3 = 25. the parameters of the bažant-najjar model and the surface factor were identified using a brute-force method implemented in a computer program. the repeatedly run analysis used the values with the ranges and stepping summarized in table 1 which makes altogether 31 360 combinations. the accuracy of the fit was assessed by means of a weighted error which considered the drying period on the log scale. the reference set of parameters which gave the best fit of with experimental data is listed in table 1 (see solid 25 pavel horák, petr havlásek acta polytechnica ctu proceedings 0 100 200 300 400 500 600 700 800 0.1 1 10 100 1000 10000 t0 = 8 d, henv = 60% calibration: experimenta xi al s hr in ka ge [1 06 ] duration of drying [day] ref exp 2 3 4 5 6 7 8 9 10 figure 1. best fits of axial shrinkage measured by bryant and vadhanavikkit [3] on 150 mm thick slabs, t0 = 8 days, henv= 60%. red line in figs. 1 and 2). this set was later used to define the reference behavior under a modified setup (i.e. different size or boundary conditions). in the parametric study the analysis starts at the age of 28 days and the ambient relative humidity is set to henv= 50% with the intention to minimize the effect of aging and to comply with the standard laboratory conditions. furthermore, as demonstrated in the following section, it was attempted to narrow the identified set of parameters by incorporating additional reference data: moisture loss (assuming k = 100 kg/m3), response to drying at different levels of relative humidity (30% and 70%) and shrinkage of a symmetrically drying slab with different thickness (75 mm and 300 mm). moreover, a model of 75 mm thick slab drying from one surface only is used to assess the influence of the identified diffusivity functions on the structural behavior. this model has the same effective thickness (and thus drying kinetics) but turns out to be significantly more sensitive to the shape of the diffusivity function. c1 α0 hc n f [mm2/day] [-] [-] [-] [mm/day] min. 35 0.02 0.68 6 1.1 max. 80 0.23 0.80 20 2.3 step 5 0.03 0.02 2 0.2 ref. 60 0.05 0.74 0.06 1.5 table 1. parameters of the transport problem: ranges, stepping for identification and the reference set. 3. results and discussion the solution obtained with the reference set of material parameters (tab. 1) is in the following figures shown consistently in thick red lines. as demonstrated in fig. 1, the axial shrinkage measured by 0 10 20 30 40 50 60 70 80 0.4 0.5 0.6 0.7 0.8 0.9 1 t0 = 8 d, henv = 60% calibration: experimentd i� us iv ity [m m 2 / da y] relative humidity [-] ref, f = 1.5 2, f = 1.5 3, f = 1.7 4, f = 1.7 5, f = 1.5 6, f = 1.5 7, f = 1.5 8, f = 1.3 9, f = 1.9 10, f = 1.7 henv figure 2. dependence of diffusivity on relative humidity for the 10 best fits in fig. 1. in the legend f is the surface factor [mm/day]. -1000 0 1000 2000 3000 4000 5000 6000 0.01 0.1 1 10 100 1000 10000 t0 = 8 d, henv = 60% calibration: experiment c ur va tu re [1 06 / m ] duration of drying [day] ref 2 3 4 5 6 7 8 9 10 figure 3. prediction of curvature of 75 mm thick slab exposed to drying from one side only at henv= 50% which started at the age t0 = 28 days. the line types and the material parameters correspond to figs. 1 and 2. bryant on prism 150 × 150 mm drying from two opposite surfaces is with very similar accuracy captured by numerous combinations of material parameters of the diffusivity function. figure 1 presents only the 10 best fits, the respective functions are shown in fig. 2 and the legend lists the values of the surface factor f [mm/day]. from the fits, it can be concluded that the evolution of axial shrinkage is almost insensitive to the diffusivity at high relative humidity (h > 0.8) where the values span from 45 to 75 mm2/day, but becomes very sensitive to the diffusivity in the region just above the ambient relative humidity henvwhere the functions form a narrow band. the response of a slab with the same effective thickness exposed to drying from one side only at standard laboratory conditions (t0 = 28 days, henv= 50%) is for the best 10 identified combinations of diffusivity shown in fig. 3. in contrast to symmetric drying, the computed response for asymmetric drying be26 vol. 30/2021 modeling of concrete subjected to symmetric and asymmetric drying 0 10 20 30 40 50 60 70 80 0.4 0.5 0.6 0.7 0.8 0.9 1 t0 = 28 d, henv = 50% calibration: simulationd i� us iv ity [m m 2 / da y] relative humidity [-] ref, f = 1.5 2, f = 1.5 3, f = 1.5 4, f = 1.5 5, f = 1.5 6, f = 1.5 7, f = 1.5 8, f = 1.5 9, f = 1.7 10, f = 1.5 henv figure 4. diffusivity functions giving the best agreement with the axial shrinkage from the simulation of symmetrically drying concrete slab (d = 150 mm, henv= 50%, t0 = 28 days) computed with the reference set of material parameters. -1000 0 1000 2000 3000 4000 5000 6000 0.01 0.1 1 10 100 1000 10000 t0 = 28 d, henv = 50% calibration: simulation c ur va tu re [1 06 / m ] duration of drying [day] ref 2 3 4 5 6 7 8 9 10 figure 5. prediction of curvature of 75 mm thick slab exposed to drying from one side only at henv= 50% and at the age t0 = 28 days. the line types and the material parameters correspond to fig. 4. comes surprisingly different. in all presented cases, the peak curvature is reached after ≈ 20 days of drying, and the individual solutions differ by up to 20% of the reference solution. in several cases the curvature exhibits a second local maximum one order of magnitude later (≈ 200 days). the value at the first peak is tightly connected with the diffusivity at high relative humidity, the higher this diffusivity, the lower the curvature and vice versa (compare e.g. the dashed blue and black lines #8 and #9). the higher this diffusivity the lower is the gradient of relative humidity between the sealed and drying surface of the cross-section and thus the lower is the resulting curvature. the second peak appears when the diffusivity function has a steep transition between high and low relative humidity which is controlled by parameter n in (2). the preceding observations are confirmed in fig 0 200 400 600 800 1000 1200 1400 0.1 1 10 100 1000 10000 t0 = 28 d calibration: simulation henv = 50% henv = 30% henv = 70% henv = 30, 50, 70% a xi al s hr in ka ge [1 06 ] duration of drying [day] ref 2 3 4 5 6 7 8 9 10 figure 6. best fits of (reference) axial shrinkage caused by symmetric drying at three levels of ambient relative humidity henv= 30, 50, 70% (d = 150 mm, t0 = 28 days). 0 10 20 30 40 50 60 70 80 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 t0 = 28 d calibration: simulation henv = 30, 50, 70% d i� us iv ity [m m 2 / da y] relative humidity [-] ref, f = 1.5 2, f = 1.5 3, f = 1.7 4, f = 1.3 5, f = 1.3 6, f = 1.7 7, f = 1.5 8, f = 1.3 9, f = 1.5 10, f = 1.7 figure 7. diffusivity functions giving the best agreement with the axial shrinkage from the reference simulations (symmetric drying at three levels of ambient relative humidity henv= 30, 50, 70%, d = 150 mm, t0 = 28 days). ures 4 and 5 where the calibration was done on the data from simulation with reference parameters but at “laboratory conditions” (t0 = 28 days, henv= 50%). the identified diffusivity functions are very close to the reference case in the region from henv to henv+ 10%–15%, while beyond this limit the functions are distinctively different which has an impact on the evolution of curvature. incorporating weight loss does not narrow the scatter of the identified diffusivity functions (figures are not presented here). from the findings presented above it is natural to expect that if the separate measurements of axial shrinkage are performed at different levels of relative humidity, then the identified diffusivity functions will be closer to each other in the relative humidity slightly above the prescribed levels. to validate this statement, 2 additional values of relative humidity 27 pavel horák, petr havlásek acta polytechnica ctu proceedings -1000 0 1000 2000 3000 4000 5000 6000 7000 0.01 0.1 1 10 100 1000 10000 t0 = 28 d, henv = 30, 50, 70% calibration: simulation c ur va tu re [1 06 / m ] duration of drying [day] ref 2 3 4 5 6 7 8 9 10 figure 8. prediction of curvature of 75 mm thick slab exposed to drying from one side only at henv= 50% which started at t0 = 28 days. the line types and the material parameters correspond to identification at 3 different levels of henv(figs. 6 and 7). 0 200 400 600 800 1000 0.1 1 10 100 1000 10000 t0 = 28 d, henv = 50% calibration: simulation 150 mm 75 mm 300 mm a xi al s hr in ka ge [1 06 ] duration of drying [day] ref 2 3 4 5 6 7 8 9 10 figure 9. best fits of (reference) axial shrinkage of 75 mm, 150 mm, and 300 mm thick slabs exposed to symmetric drying at henv= 50% (t0 = 28 days). 0 10 20 30 40 50 60 70 80 0.4 0.5 0.6 0.7 0.8 0.9 1 t0 = 28 d, henv = 50% calibration: simulation d i� us iv ity [m m 2 / da y] relative humidity [-] ref, f = 1.5 5, f = 1.5 6, f = 1.5 8, f = 1.5 figure 10. diffusivity functions giving the best agreement with the reference simulations of axial shrinkage of symmetrically drying slabs with various thicknesses (henv= 50%, t0 = 28 days, fig. 9). 0 10 20 30 40 50 60 70 80 0.4 0.5 0.6 0.7 0.8 0.9 1 t0 = 28 d, henv = 50% calibration: reverse d i� us iv ity [m m 2 / da y] relative humidity [-] ref, f = 1.5 2, f = 1.5 3, f = 1.5 4, f = 1.5 5, f = 1.5 6, f = 1.5 7, f = 1.7 8, f = 1.5 9, f = 1.5 10, f = 1.5 figure 11. diffusivity functions calibrated inversely on the evolution of curvature of 75 mm thick slab drying from one side (henv= 50, t0 = 28 days) obtained from the simulation with the reference set of material parameters. were added and the analysis was done at 30%, 50% and 70%. the 10 best fits are almost indistinguishable from the reference solution of axial shrinkage (see fig. 6) and indeed, the identified diffusivity functions form a narrow band which widens with increasing relative humidity (fig. 7). this has a direct impact on the computed curvature shown in fig. 8 where the scatter in the first peak persists (caused by high scatter of the maximum diffusivity), but beyond 100 days of drying the agreement between the solutions becomes significantly better. next, it was attempted to identify the diffusivity function based on the shrinkage evolution of specimens (i.e. models) with different effective sizes. the reference thickness d = 150 mm was doubled to 300 mm and reduced to 75 mm. the 10 best fits match the solution with reference parameters perfectly (fig. 9), yet the diffusivity functions show considerable scatter, see fig. 10. also it needs to be emphasized that such approach can turn out to be very demanding on the experimental duration since the evolution scales with d2. to summarize, and as illustrated in figs. 12 and 13 it turned out that the shrinkage of symmetrically drying specimen is significantly less sensitive to the diffusivity function compared to the curvature produced by non-symmetric drying. for this reason the diffusivity identified from axial shrinkage should be used with caution and to members with similar size and boundary conditions. on the other hand the diffusivity identified from the evolution of curvature (see fig. 11) exhibits significantly less scatter and so does the back-calculated axial shrinkage which matches the reference case perfectly (not presented here). 28 vol. 30/2021 modeling of concrete subjected to symmetric and asymmetric drying 0 10 20 30 40 50 60 70 80 0.4 0.5 0.6 0.7 0.8 0.9 1 t0 = 28 d, henv = 50% c1 60, alpha0 0.05, hc 0.74, n 6, f 1.5 c1 75, alpha0 0.10, hc 0.80, n 16, f 1.1 d i� us iv ity [m m 2 / da y] relative humidity [-] reference, f = 1.5 modi�ed, f = 1.1 henv figure 12. diffusivity function with the “reference” and “modified” set of material parameters. 4. conclusions this paper investigated drying shrinkage of concrete slabs exposed to symmetric and asymmetric drying at constant ambient relative humidity. the problem was approached by means one-way coupled finite element simulations in program oofem. the moisture transport was described using the the bažant-najjar model while the time-dependent behavior of concrete by the modified mps model. the emphasis was on put on identification of material parameters of the model for moisture transport and their uniqueness. the following conclusions can be drawn from the analysis: • identification of parameters based the evolution of axial shrinkage of symmetrically drying concrete does not yield unique set of material parameters. the reference data could be reproduced with various sets of parameters with accuracy similar or better than the typical experimental scatter. • incorporating the shrinkage data measured at different levels of relative humidity helps to narrow the identified diffusivity functions. this is not the case of the data on moisture loss or shrinkage of specimens with different size. • the behavior of concrete drying non-symmetrically is considerably more sensitive to the values of material parameters. the parameters identified from the evolution of curvature give more consistent diffusivity functions and result into almost identical axial shrinkage. • the presence and the character of the second maximum in the evolution of curvature is most sensitive to the exponent n which defines the steepness of the transition between small and high diffusivity. • the numerical simulations of concrete drying from one side only exhibit very little sensitivity to concrete creep (fig. 13). -200 0 200 400 600 800 1000 1200 0.01 0.1 1 10 100 1000 10000 -1000 0 1000 2000 3000 4000 5000 6000 t0 = 28 d, henv = 50% a xi al s hr in ka ge [1 06 ] c ur va tu re [1 06 / m ] duration of drying [day] reference modi�ed figure 13. evolution of axial shrinkage and curvature obtained with the “reference” and “modified” set of material parameters of the diffusivity function (fig. 12) and d = 150 mm, t0 = 28 days, henv= 50%. thin lines correspond to the behavior without viscoelastic effects. acknowledgements the authors gratefully acknowledge financial support from the grant agency of the czech technical university in prague, project number sgs20/038/ohk1/1t/11, and from the czech science foundation (ga čr), project number 19-20666s. references [1] v. šmilauer, p. havlásek, t. gasch, et al. hygro-mechanical modeling of restrained ring test: cost tu1404 benchmark. construction and building materials 229:116543, 2019. doi:https://doi.org/10.1016/j.conbuildmat.2019.07.269. [2] b. patzák. oofem home page, 2000. http://www.oofem.org. [3] a. h. bryant, c. vadhanavikkit. creep, shrinkage-size, and age at loading effects. aci materials journal 84:117–123, 1987. [4] z. p. bažant, l. j. najjar. nonlinear water diffusion in nonsaturated concrete. materials and structures 5:3–20, 1972. doi:10.1007/bf02479073. [5] fédération internationale du béton. model code 2010. no. vol. 65 in fib bulletin. international federation for structural concrete (fib), 2012. [6] z. bažant, p. havlásek, m. jirásek. microprestress-solidification theory: modeling of size effect on drying creep. in n. bicanic, h. mang, g. meschke, r. de borst (eds.), computational modelling of concrete structures, pp. 749–758. crc press/balkema, eh leiden, the netherlands, 2014. [7] z. p. bažant, a. p. hauggaard, s. baweja, f. j. ulm. microprestress solidification theory for concrete creep. i: aging and drying effects. journal of engineering mechanics 123:1188–1194, 1997. 29 https://doi.org/https://doi.org/10.1016/j.conbuildmat.2019.07.269 http://www.oofem.org https://doi.org/10.1007/bf02479073 pavel horák, petr havlásek acta polytechnica ctu proceedings [8] z. bažant, s. baweja. creep and shrinkage prediction model for analysis and design of concrete structures: model b3. adam neville symposium: creep and shrinkage structural design effects 2000. [9] p. havlásek. creep and shrinkage of concrete subjected to variable environmental conditions, phd. thesis. czech technical university in prague, 2014. 30 https://doi.org/10.14311/app.2022.33.0539 acta polytechnica ctu proceedings 33:539–545, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague quantitative analysis of the strength generating c-s-h-phase in concrete by ir-spectroscopy frank michael schmidt-döhla, ∗, david schulenberga, franziska tralowa, jürgen neubauerb, julian johannes wolfb, dominique ectorsb a hamburg university of technology, tuhh, institute of materials, physics and chemistry of buildings, 21071 hamburg, germany b friedrich-alexander university erlangen-nürnberg, fau, geozentrum nordbayern, schlossgarten 5a, 91054 erlangen, germany ∗ corresponding author: schmidt-doehl@tuhh.de abstract. the c-s-h-phase is the most important strength generating phase in concrete and other cementitious materials. the analysis of c-s-h is therefore an important instrument of innovations in the field of concrete and its durability and sustainability. the quantification and insights in c-s-h are hindered by the predominantly amorphous structure of c-s-h. only the time consuming and expensive solid state nuclear magnetic resonance spectroscopy (nmr) gives a chance to get results, but is restricted to model substances. a new technique, based on the cheap and widespread infrared-spectroscopy (ir) was developed. the quantitative analysis is based on the silicon content in the different structural units of c-s-h and a calibration with natural and synthetic materials with known silicon content in these units. the technique allows to investigate specimens from real structures. specimens with quartz or limestone aggregate can be analyzed directly. in other cases the aggregates must be separated, for example with heavy liquid separation. tests with different mortars showed a good correspondence of measured and expected values of c-s-h concentration. keywords: calcium-silicate-hydrate (c-s-h), quantitative analysis, spectroscopy. 1. introduction undoubtedly the c-s-h-phase is one of the most important chemical substances in the world. it is generated by the hydration of c3s and c2s, puzzolanic reactions and reactions of latent hydraulic components. it is the strength building phase in most cementitious material. if researchers want to improve mineral building materials, and as a consequence improve sustainability of such materials, it is therefore of great importance to quantitatively analyze c-s-h and its variable, actual structure. also in the case of corrosion and durability investigations, where the cs-h is attacked and the concentration is decreased, it will be very helpful to quantitatively analyze c-s-h. still by the extensive amorphous character of c-s-h in x-ray diffraction experiments and other possible amorphous components in the material it is not possible to quantitatively analyze c-s-h at the moment in a general manner. only model substances can be analyzed quantitatively, for example hydrated alite or hydrated white cement with 29si magic angle spinning solid state nuclear magnetic resonance, which is a very expensive and time consuming technique. also special systems can be analyzed with quantitative x-ray diffraction by calculating the c-s-h from the changes of the crystalline species. up to now it is not possible to take a specimen from a real structure and make a quantitative analysis (qxrd). this was the motivation for the development of the method. 2. basic information the method is based on attenuated total reflection (atr) fourier transform (ft) infrared spectroscopy (ir), with a single reflection diamond atr. the generated evanescent wave at the interface is attenuated by the absorption in the specimen, based on the excitation of vibrations in functional groups with a change of the dipole moment. nearly no scattering effects are generated by this technique and the penetration of the evanescent wave into the specimen is less than 5 microns. a quantitative analysis with the atr-technique is based on lambert-beers law, like in transmission. that means the absorbance is directly proportional to the concentration, except very high concentrations [1]. to our experience it is necessary to dilute the calibration substances at least to 50 % to make sure that the calibration curve is in the linear range. the basic structure of c-s-h consists of short chains of sio4-tetrahedrons. the so called q1 structural units are the ends of the chains, the q2 structural units are in the middle. the chain length is connected with the ratio q1/q2. infrared spectroscopy produces direct signals from c-s-h, generated by the 539 https://doi.org/10.14311/app.2022.33.0539 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en f. m. schmidt-döhl, d. schulenberg, f. tralow et al. acta polytechnica ctu proceedings q1 and q2 structural units. they can be used to develop a method for the quantitative analysis of c-s-h. the bands are weak and broad and there are overlapping signals from other phases. and, of course, a quantitative analysis needs a calibration. these difficulties must be overcome. for the development of the method different portland cements, blastfurnace slags, flyashes, silica fume, lime stone powders, quartz powder, hydrated lime, mixtures of these components, and hydrated mortars made of these components were used. the components were characterized by polarized light microscopy, x-ray diffraction, x-ray fluorescence and infrared spectroscopy. the measurements were performed with a perkin elmer spectrum two with universal attenuated total reflectance accessory and the perkin elmer software spectrum. peak fitting was performed with the program peakfit of systat software, erkrath, germany. 3. method for the quantitative analysis of c-s-h in cementitious materials by atr-ir-spectroscopy a detailed inspection instruction is available from the corresponding author (in german). 1. at first a qualitative phase analysis of the specimen is needed. due to the overlapping of peaks, it is important to know the phase assemblage in the sample. 2. the specimen is crushed to a grain size of some millimeters and dried. the drying must be restricted to a minimum of 11 % r.h., because a stronger drying can affect the c-s-h phase. then the specimen has to be ground to a powder. with an atr crystal area of four square millimeters it is necessary to grind to a grain size of about thirty to forty micrometers to get sufficient grain statistics of several thousand grains. dry grinding is possible, but it is necessary to avoid temperatures above 70 ◦c. the powder must kept dry and without carbonation. 3. other minerals than quartz and calcite in the aggregates must be separated from the specimen, because of possible peak overlapping. this separation starts with a careful crushing in step 2, where big aggregates can be separated by hand. the further separation of aggregates can be done with heavy liquid separation with diiodomethane/acetone mixture with a density of 2.4 g/cm3. [2] the separation must be done quantitatively. this technique does not affect the infrared spectra of c-s-h. the very common minerals quartz and calcite need not to be separated. therefore it is possible to produce laboratory concrete specimens, which can be analyzed without a heavy liquid separation. the aggregates from concrete specimens from buildings must be separated normally, because of the many different aggregate types inside. 4. the method needs an internal standard to correct differences between the density of the specimens on the atr-crystal and changes in the intensity of the light source of the spectrometer. the scattering of the c-s-h signals is decreased by working with a relative absorbance. we found potassium ferricyanide, k3[fe(cn)6] to be sufficient as an internal standard for this method. it must be grinded wet in water free ethanol to a grain size of about 20 micrometers without remarkable increase of temperature. any contact with acid, chromium trioxide, oxidizing agents, fluorine, ammonia, nitrates and nitrites must be avoided. because of its solubility in water it is usable only with dry specimens. the standard shows two peaks at 2117 and 2120 cm−1 and a weak band at 500 cm−1. it must be kept dry and guarded from light. 5. the specimen and the internal standard are mixed in a 60 to 40 relation by weight. this mixing must be done wet in a little amount of ethanol, to get a homogenous mixture. the precision of weighting is critical. ethanol must completely evaporate before the measurement. 6. the specimen, mixed with the internal standard is measured with a resolution of 2 cm−1. the atr crystal must be completely covered with powder and the pressure on the powder must be controlled so that the spectrum is not changed with increasing pressure. at least 5 measurements of the material are recommended. 7. the spectrum is converted into absorbance, and an atr correction is performed, assuming ideal contact of powder to the crystal. the atr correction is a mathematical transformation, converting the spectrum into a spectrum comparable with a transmission spectrum. 8. now it is necessary to eliminate interfering signals from other phases, known by step 1. interfering signals can be separated within the necessary peak fit (see the following step). but in the case of components with very broad ir-signals, for example flyash or blastfurnace slag, it is helpful to subtract a typical spectrum of such a component, measured with the same conditions. often it is not possible or recommended to subtract all of the intensities of the interfering substances. the residual intensities will be eliminated by the peak fit, remembering the peaks of the interfering substances. 9. in the next step a peak fit in the area of the c-s-h peaks and of the internal standard is performed. during this peak fit, also the remaining interfering signals from other phases and the background must be considered. we prefer a linear background between 1850 and 2250 cm−1 for the internal standard, and a background with logarithm function 540 vol. 33/2022 quantitative analysis of c-s-h figure 1. calibration curves for the q1 (left) and q2 (right) structural units in c-s-h. in the range of the c-s-h bands, leaned on the minimum points in this range. it is absolutely necessary to restrict the peaks to that, existing in the different phases of the specimen. it is necessary to make this fit with experience on the technique, knowledge of the phases in the specimen and their typical peaks. the absorbance is measured as peak area (voigt-function). 10.then, with the absorbance aq of the q1 and q2 signals of c-s-h and the absorbance of the internal standard ai , a relative absorbance arel = aq/ai of q1 and q2 is calculated. 11.with the help of calibration functions the relative absorbance is changed into a concentration of silicon in the q1 and q2 structural elements of c-s-h in the specimen. 4. calibration with crystalline substances unfortunately there is no c-s-h standard existing, were the concentration of si in the q1 and q2 structural elements is known. therefore we tried to use minerals containing q1 or q2 structural elements as calibration substances to develop the calibration curves for c-s-h. for the calibration of the silicon in the q2 structural units we used the natural minerals cordierite, wollastonite and beryl. for the calibration of the silicon in the q1 structural units the synthesized minerals åkermanite and gehlenite were used. all calibration standards were characterized by x-ray fluorescence, x-ray powder diffraction and ir-spectroscopy. the calibration substances were diluted with fly ash, blastfurnace slag, lime powder, unhydrated portland cement, hydrated lime and silica fume to get specimens with different content of q1 resp. q2 groups. these specimens were measured and the data processed as described in chapter 3. because of the lack of space no more details of this calibration can be presented here. figure 1 shows the calibration curves for silicon in the q1 and the q2 structural units in c-s-h with the data of all calibration substances, diluted by different substances common in cementitious materials, named at the beginning of this chapter, and analyzed by three different people. these calibration curves refer to dry specimens. it can be seen in these diagrams, that although very different calibration substances were used, with different dilution components, and different people performing these measurements, they all fit to the same calibration curves, within an expectable scatter. 5. calibration using in addition fully hydrated activated alite plausibility tests were performed by analyzing fully hydrated activated alite, different hydrated binders, and mortars (with binder enrichment) with wellknown composition. using the calibration described in chapter 4 the tests systematically result in too high total silicon concentrations in the specimens. so the calibration and the results were checked very carefully again. in addition the measured chain lengths are in the range of data reported in the literature, based on 29si mas nmr measurements. this indicates that the measured q1/q2-ratio is correct, assuming that the nmr-data are right. so it seems that the calibration of the q1/q2-ratio by measuring crystalline minerals is possible. but the total si-content must be calibrated with c-s-h specimens with a known sicontent. we used therefore fully hydrated activated alite, described in chapter 6 for the calibration of the total si content in c-s-h. with this technique we got the following calibration functions. csi in q1 = arel, q1 ! (0.163 + 0.122) (1) csi in q2 = arel, q2 ! (1.017 + 0.763) (2) • csi concentration of si in the q1 or q2 structural units of c-s-h respectively • arel relative absorbance of the q1 or q2 peaks respectively, described in chapter 3 541 f. m. schmidt-döhl, d. schulenberg, f. tralow et al. acta polytechnica ctu proceedings age 24 hours 7 days 6 months wt.-% si in q2 in specimen 3.4 ± 0.2 4.0 ± 0.2 3.9 ± 0.2 wt.-% si in q1 in specimen 2.0 ± 0.1 2.1 ± 0.2 1.9 ± 0.2 sum of two lines above 5.4 ± 0.3 6.1 ± 0.3 5.8 ± 0.4 mean chain length 5.4 ± 0.2 5.8 ± 0.4 6.1 ± 0.3 table 1. plausibility tests with fully hydrated activated alite. age 24 hours 7 days 6 months measured weight percent c-s-h (as c1.7sh2.6) 34.8± 1.2 36.4± 0.1 37.4± 1.0 measured weight percent portlandite 20.7± 0.4 20.5± 0.5 20.0± 0.5 table 2. xrd measurements with fully hydrated activated alite. the first numeric value in the equations originates from the calibration with the crystalline substances (see chapter 4). the second numeric value originates from the correction of the total si-content, measured with fully hydrated activated alite. the measurement of the same hydrated cement stone in different laboratories showed, that it will be necessary to make an instrument specific calibration. it is one result of the research that this calibration can be reduced to a one-point-calibration of the second numeric value in equations (1) and (2), together with the proof that without c-s-h in unhydrated cementitious binder the measured absorbance of q1 and q2 signals will be zero. the authors can deliver fully hydrated activated alite to perform such a calibration. the typical quantitation limit of silicon in the q1 or q2 structural unit of 0.3 wt.-% for a number of measurements of 5 and a significance level of 5 % in most cases is sufficient not only for hydrated binders, but also for mortar and concrete. otherwise an enrichment of the binder and/or an increased number of measurements are necessary. 6. plausibility tests with fully hydrated alite activated alite (71.7 wt.-% cao, 25.9 wt.-% sio2, 1.8 wt.-% mgo, 0.6 wt.-% al2o3) was hydrated with a water cement ratio of 1.1 at 23 ◦c in closed containers without moisture loss or uptake from the environment. the total concentration of silicon in this system is very well known. it amounts to 5.72 wt.-%. xrd measurements after 24 h reaction time with water showed no c3s anymore. all of the silicon must therefore be incorporated in the reaction product cs-h. but the detailed structure of the c-s-h is unknown and a function of age. the hydrated system was analyzed by atr ft ir spectroscopy after 24 h, 7 days and 6 months reaction time. differing from the method described in chapter 3 the specimen was not dried until the mixing with the internal standard. additional measurements with thermal analysis proofed, that the moisture loss until mixing with the internal standard was maximum 0.9 wt.-%, increasing the silicon content in the system to maximum 5.76 wt.-%. additional water loss after the mixing with the internal standard do not affect the analysis results. the results are shown in table 1. it is visible, that the mean chain length is slightly increasing with time. the mean chain length is near chain lengths based on nmr-data in the scientific literature [3–6]. the total silicon in c-s-h of course is near the theoretical value of 5.72 wt.-%, because the specimens were used for the calibration of this parameter. in addition the material was analyzed independently by x-ray diffraction with the method of bergold et al. [7] at the same time. assuming a c-s-h composition of c1.7sh2.6, 41.2 wt.-% c-s-h and 20.1 wt.-% portlandite can be expected. the results of the measurements are shown in table 2. it can be seen, that the c-s-h content is lower than the expected one. this was also the result of bergold et al. [7] and is an indication that a part of the c-s-h is totally x-ray amorphous. this may be the dimer molecules. the xrd detectable amount of c-s-h is increasing with time. figure 2 shows the spectrum of the fully hydrated activated alite with age 6 months in the range of the si-o stretching vibrations in c-s-h. it shows 4 peaks resulting from the q2 structural unit at 910, 954, 995 and 1064 cm−1. at the age of 7 and 1 days only the two strongest peaks in the middle can be observed. the q1 structural unit shows two peaks at 793 and 820 cm−1. at 1 day only the strongest peak above 800 cm−1 can be observed. at 7 days the weak splitting is observable. the full spectrum additionally contains the common bands resulting from portlandite, the internal standard, water, carbonate, a peak at 1325 cm−1 perhaps a si-oh bending vibration in cs-h [8], a weak peak at 655 cm−1 and several peaks in the area 450 545 cm−1, perhaps resulting from deformations of the tetrahedrons or the tetrahedron chain. 542 vol. 33/2022 quantitative analysis of c-s-h figure 2. atr ir spectrum of fully hydrated activated alite with age 6 months in the range of the most important bands of c-s-h including the peak fit. background (logarithm function) subtracted before, leaned on the minima at the end of the displayed range. note that the signal of the q1 structural element seems to be more sharp and restricted to lower wave numbers than in the case of a commercial cement stone (see figure 3). cement cem i42,5 r cem i 42,5 r-sr 0 cem i 52,5 r cem i 52,5 r cem i 52,5 r cement g 680 680 780 325 590 water g 237.82 267.72 232.88 194.7 353.44 silica fume elkem 940 u g − − − − 65.52 water-cement ratio 0.35 0.4 0.3 0.6 0.6 water-binder ratio − − − − 0.54 degree of hydration c3s % 81.3 93.7 73.7 99.4 95.6 degree of hydration c2s % 18.9 35.5 7.8 43.9 35.0 wt.-% si in q2 in specimen 3.6 ± 0.2 3.8 ± 0.2 3.7 ± 0.2 4.6 ± 0.4 5.5 ± 0.4 wt.-% si in q1 in specimen 1.8 ± 0.3 1.3 ± 0.2 1.9 ± 0.2 1.9 ± 0.2 2.4 ± 0.2 sum of two lines above 5.4 ± 0.3 5.1 ± 0.3 5.6 ± 0.2 6.5 ± 0.3 7.9 ± 0.5 expected value (see text) 5.3 6.4 4.9 7.4 10.0 mean chain length (calc.) 6.0 ± 0.6 7.9 ± 1.1 6.0 ± 0.6 6.9 ± 0.6 6.6 ± 0.4 table 3. plausibility tests with hydrated binders. 7. plausibility tests with hydrated binders and mortars using the phase analysis of the cements measured by the rietveld method, the degree of hydration of c2s and c3s measured by the g-factor method [9], an assumed degree of hydration of 100 % for silica fume, the composition of the mixtures, and the drying conditions, the amount of silicon in the q1 and q2 structural elements of c-s-h was calculated in weight percent of the specimens and compared with the experimentally determined value for several hydrated binders. all specimens were stored in closed containers at 23 ± 2 ◦c and above 95 % r.h. until drying. the age before drying was at least 84 days. peaks at about 850 and 900 cm−1 were treated as residual cement phases in this case. table 3 summarizes the tests and their results. it can be seen, that the maximum deviation from the expected total silicon content of the specimens cs-h is about 20 %. positive and negative deviations could be observed. the specimen with the w/c-ratio 0.6 contents more c-s-h as the specimen with the same cement and the w/c-ratio 0.3, as a result of the higher degrees of hydration. the increase of cs-h content by the addition of silica fume is very well visible. the mean chain length of the hydrated binders amounts to 6.0 7.9 and is near chain lengths based on nmr-data in the scientific literature [3–6]. in the same way the amount of silicon in c-s-h of several hydrated mortar binders was analyzed. in this case, the total water content of the specimens was determined by thermal analysis as loss on ignition. after production all mortar specimens were 543 f. m. schmidt-döhl, d. schulenberg, f. tralow et al. acta polytechnica ctu proceedings aggregate type granite, sand basalt, sand limestone, sand quartz sand granite, basalt, limestone g 962.42 962.42 962.42 0 sand, mainly quartz g 641.62 641.62 641.62 1738.07 cement type cem i 52,5 r cem i 52,5 r cem i 52,5 r cem i 52,5 r cement g 535 535 535 580 water-cement ratio 0.5 0.5 0.5 0.55 heavy liquid separation yes yes yes no degree of hydration c3s % 100 100 100 100 degree of hydration c2s % 98.1 98.7 97.2 100 residual phases of aggregate wt.-% quartz 1.8 quartz 1.8 calcite 10.4quartz 1.3 quartz 12.4 wt.-% si in q2 in specimen 3.3 ± 0.3 3.3 ± 0.3 2.9 ± 0.1 3.8 ± 0.1 wt.-% si in q1 in specimen 2.5 ± 0.3 2.3 ± 0.3 2.0 ± 0.1 2.0 ± 0.3 sum of two lines above 5.8 ± 0.6 5.6 ± 0.4 4.9 ± 0.2 5.8 ± 0.3 expected value (see text) 7.6 7.7 6.9 6.9 mean chain length (calc.) 4.6 ± 0.2 4.9 ± 0.4 4.8 ± 0.1 5.8 ± 0.6 table 4. plausibility tests with binders of hydrated mortars. figure 3. atr-ir spectrum including the peak fit of a mortar with quartz aggregate and enrichment of the binder by grinding and sieving, showing the peaks of the q1 and the q2 structural units in the c-s-h. background (logarithm function) subtracted before, leaned on the minima at the end of the displayed range. stored in closed containers at 23 ± 2 ◦c and above 95 % r.h. until drying. the age before drying was at least 1.5 years (mortar with only quartz: about 0.5 years). the aggregates of the mortar specimens were then removed by two steps: 1) repeated grinding and sieving, removing the coarse grains, 2) heavy liquid separation of the fines with diiodomethane/acetone mixture with a density of 2.4 g/cm3 (only mortars with basalt, granite and limestone). table 4 summarizes the tests and their results. all specimens contain a small amount of residual quartz, probably from the quartz sand. the mortar with only quartz aggregate contains a much higher value, because a heavy liquid separation was not performed in this case. all specimens contain about 0.5 wt.-% calcite, probably by carbonation. the specimen with limestone aggregate contains a much higher value. this indicates that the limestone aggregate cannot be removed well by the heavy liquid separation. but quartz and calcite can be treated during the evaluation of the spectra. no typical minerals from the basalt and the granite could be observed after heavy liquid separation. astonishing is the very high degree of hydration of the c2s. this cannot be an effect of the heavy liquid separation, because also the mortar with quartz and without heavy liquid separation has a degree of hydration of 100 %. figure 3 shows the peak fit of a specimen with quartz sand. because of the high degree of hydration no residual cement phases were considered. in this case all measured si-contents in the specimens c-s-h are lower than the expected values, and the deviation is greater (maximum 29 %). 544 vol. 33/2022 quantitative analysis of c-s-h because of the broader q1-peak used in these peak fits (compare figure 2 and figure 3), the calculated mean chain length is smaller than in the alite and the hydrated binder evaluation. 8. conclusions it can be concluded, that atr-infrared spectroscopy is a very helpful and powerful technique, also in a quantitative manner. infrared spectroscopy is used for the analysis of cementitious materials for a long time, but its potential is not exhausted. also amorphous species are producing signals that can be observed directly. this feature was used to develop a quantitative analysis method for the structurally only weakly ordered c-s-h in cementitious materials. with this technique it is possible to analyze specimens from real life, not only model materials. cementitious materials with quartz or calcite aggregates can be analyzed directly, although an enrichment of the binder will facilitate the analysis. other aggregates must be separated, for example with the help of heavy liquid separation. in addition infrared spectroscopy is a very fast and cheap technique that is common in most laboratories, active in the field of construction chemistry. the described method measures the silicon content in the q1 and q2 structural units in the c-s-h. crystalline minerals containing q1 or q2 structural units were used to calibrate the q1/q2 ratio. but it was not possible to use these crystalline materials for the total silicon content as well. the total silicon content was calibrated with the help of an activated fully hydrated alite with very well known silicon content in the c-s-h phase. plausibility tests on different hydrated binders and the binder in mortars showed that the total silicon content in c-s-h, measured with the method, is in the range of the expected values, but deviations of up to 30 percent could be observed. in particular the measurement of the small q1 absorbance is difficult. in addition also the ratio q1/q2 is measurable and therefore the chain length, which is in the range of literature data based on nmr. in spite of these difficulties researchers and manufacturers have the chance to analyze the concentration and the internal structure of c-s-h as a function of relevant parameters with this method. not only the success of improvements of the time dependent c-s-h quantity and structure, but also the effects of other measures on c-s-h and the results of decomposition of c-s-h by chemical attack can be investigated. in the future it will be helpful to investigate how aluminium and alkalies in c-s-h will affect the irspectra of c-s-h and the method. in addition it is necessary to find an alternative internal standard, that is not soluble in ethanol and water, has a good grindability and chemical stability, is not toxic and does not overlap with the interesting peaks. further activity should pay attention also on the calibration substance and the removal of interfering aggregates. acknowledgements some of the data measurement was performed by m. maaser and s.k. mehra. funding: this work was supported by the deutsche forschungsgemeinschaft [grant number 280067300]. references [1] m. milosevic. internal reflection and atr-spectroscopy, hoboken, wiley, 2012. [2] c. jehn. transportprozesse und chemische reaktionen in silikatischen gesteinskörnern, hamburg: dissertation hamburg-harburg university of technology, 2015. [3] f. battocchio, p. j. m. monteiro, h.-r. wenk. rietveld refinement of the structures of 1.0 c-s-h and 1.5 c-s-h. cement and concrete research 42(11):1534-48, 2012. https: //doi.org/10.1016/j.cemconres.2012.07.005. [4] e. gallucci, x. zhang, k. l. scrivener. effect of temperature on the microstructure of calcium silicate hydrate (c-s-h). cement and concrete research 53:185-95, 2013. https: //doi.org/10.1016/j.cemconres.2013.06.008. [5] i. g. richardson. the calcium silicate hydrates. cement and concrete research 38(2):137-58, 2008. https: //doi.org/10.1016/j.cemconres.2007.11.005. [6] j. skibsted, m. d. andersen, h. j. jakobsen. application of solid state nuclear magnetic resonance (nmr) in studies of portland cement-based materials, zkg int. 60:70-83, 2007. [7] s. t. bergold, f. goetz-neunhoeffer, j. neubauer. mehcanically activated alite: new insights into alite hydration. cement and concrete research 76:202-11, 2015. https: //doi.org/10.1016/j.cemconres.2015.06.005. [8] r. ylmén, u. jäglid, i. panas, et al. monitoring early hydration of cement by ex situ and in situ atr-ftir a comparative study. journal of the american ceramic society 97(11):3669-75, 2014. https://doi.org/10.1111/jace.13186. [9] d. jansen, f. goetz-neunhoeffer, c. stabler, et al. a remastered external standard method applied to the quantification of early opc hydration. cement and concrete research 41(6):602-8, 2011. https: //doi.org/10.1016/j.cemconres.2011.03.004. 545 https://doi.org/10.1016/j.cemconres.2012.07.005 https://doi.org/10.1016/j.cemconres.2013.06.008 https://doi.org/10.1016/j.cemconres.2007.11.005 https://doi.org/10.1016/j.cemconres.2015.06.005 https://doi.org/10.1111/jace.13186 https://doi.org/10.1016/j.cemconres.2011.03.004 https://doi.org/10.14311/app.2022.33.0344 acta polytechnica ctu proceedings 33:344–349, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague extent of corrosion damage for rc structures exposed to chloride-bearing environment federica lollini politecnico di milano, department of chemistry, materials and chemical engineering "giulio natta", piazza leonardo da vinci 32, italy correspondence: federica.lollini@polimi.it abstract. in industrialized countries, most of the reinforced concrete (rc) structures and infrastructures have more than 40 years, since they were built around 1960 − 1980, and were designed for a service life of 50 years. consequently, in the next years, the number of structures and infrastructures that will need to be repaired will steeply increase. several techniques, characterized by a different durability, environmental impact and economic impact, are available for the repair of a reinforced concrete structure damaged by the corrosion of reinforcement. to help the designers in the choice of the most suitable one, a preliminary assessment of the condition of the structure is essential, aimed at diagnosing the causes of deterioration, the extent of damage, i.e. the extent of corroding reinforcement, and its evolution in time. an approach, with a wide consensus, for the evaluation of the extent of corroding reinforcement is available for carbonation-induced corrosion, whilst it is still lacking for chloride-induced corrosion. in this paper two approaches for the evaluation of the extent of corroding reinforcement for a rc structure subject to chloride-induced corrosion are presented and compared, showing similar results. these approaches can be useful to properly plan a restoration intervention as well as to assess the reliability of the recently proposed model for service life design. keywords: assessment, corrosion, chloride. 1. introduction in industrialized countries, most of the reinforced concrete (rc) structures and infrastructures, were built around 1960 − 1980 and, in general, were designed to last for a period of 50 − 100 years [1]. more and more of these structures are approaching the end of their designed service life, and due to their exposure conditions, i.e. typical urban or rural environment or marine exposure, often suffer damage due to carbonationor chlorideinduced corrosion of reinforcement, such as concrete spalling or cracking or rust stains that may affect their serviceability. a proper maintenance and rehabilitation of reinforced concrete structures could extend their service life, by restoring the structural safety and preventing the future damage. this would avoid to demolish them and build new ones, leading to a lower social, environmental and economic impact, due, for instance, to a less hindrance to traffic and waste of materials [2]. several techniques are available for the rehabilitation, which vary from keeping the conditions of corrosion and serviceability under control by some form of monitoring, when the extent of damage is limited or the remaining period of use is short, to the conventional repair technique, which requires the replacement of carbonated concrete cover or chloride contaminated cover with a cementitious mortar, to electrochemical techniques or hydrophobic treatments, that can control corrosion of steel bars without requiring the removal of the original concrete cover, except the cracked concrete [3–5]. each of these strategies is characterizes by a different durability, environmental impact and economic impact, and the choice of the most suitable protection system should be a compromise among the different requirements [4]. hence, it appears essential to choose and plan the most suitable restoration intervention, a careful condition investigation that cannot be limited to a visual inspection by an expert, but should include, beside the visual inspection and the examination of design documents, field tests as well as sampling for laboratory tests, i.e. carbonation depth, chloride profile and cover depth. the inspection should be aimed at estimating if, both at the time of inspection and at the end of the designed lifetime, the reinforcement is still passive, i.e. corrosion has not initiated since carbonation or a critical amount of chlorides has not reached the steel surface, or the reinforcement is corroding but the propagation is in the early stages, e.g. concrete cover is not cracked and reduction in cross section of rebars is negligible [6]. since concrete quality, mainly the permeability, nature and intensity of cracks, and concrete cover are spatially variable over the concrete surface, due to different concrete batches and the heterogeneity of workmanship, the corrosion conditions will vary accordingly [7–9]. an evaluation of the extent of corroding reinforcement, i.e. of the amount of steel bars that are no more longer passive, and its evolution in time, is, then, needed. for carbonation-induced corrosion the extent of corroding reinforcement can be assessed by comparing measured carbonation depth with measured cover depths 344 https://doi.org/10.14311/app.2022.33.0344 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 corrosion damage for rc structures d�� e�� � ��� � ��� � ��� � ��� ����������������������������� c l wk �� �p dv v� ri �f hp hq w� 3uredelolw\ pfruu � ��� � ��� � ��� � ��� � �� �� �� �� �� �� �� c & o �� �p dv v� ri �f hp hq w� 'hswk��pp� 3uriloh�� 3uriloh�� 3uriloh�� 3uriloh�� 3uriloh�� figure 1. cumulative distribution function of the critical chloride content (a) and imaginary chloride profiles measured on a rc structure (b). in the structure and in the literature a statistical approach with a wide consensus was proposed for its evaluation [10]. for chloride-induced corrosion, to evaluate the extent of corroding reinforcement, the chloride content at the depth of the bars needs to be compared to the critical chloride content, since corrosion occurs when at the bar depth a chloride content equal to the critical chloride content is reached. the variability of these parameters makes it difficult to assess the extent of corroding reinforcement and, unfortunately, an approach with a wide consensus is still lacking. this paper presents two mathematical approaches for the evaluation of the extent of corroding reinforcement for a reinforced concrete structure subjected to chloride-induced corrosion. a fictitious example is provided to compare the results of the two approaches and to discuss their real application. 2. mathematical approaches for the evaluation of the extent of corroding reinforcement 2.1. statistical approach in order to assess the extent of corroding reinforcement, rdep, of an existing structure, the critical chloride content, clth, the chloride content, ccl, and the depth of the bars, c, need to be known. the critical chloride content depends on several factors, and, due to its stochastic nature, can be properly defined only through a probability density function. according to the model code for service life design, proposed in 2006 by the international federation of concrete, fib [11], clth for ordinary carbon steel bars, can be described by a beta distribution, lower and upper limited. figure 1a shows the cumulative distribution function of the critical chloride content, assuming that the average value and the standard deviation are respectively 0.6 % and 0.15 % vs mass of cement (values of 0.2 and 2 % vs mass of cement have been assumed for the lower and upper limits). the chloride content at the different depths and the concrete cover thickness can be determined during the inspection of the rc structure, through field and laboratory tests. in order to assess rdep, and to increase its reliability and accuracy, a significant number of measurements of these parameters is required. as an example, figure 1b shows imaginary chloride profiles measured on five concrete samples taken from the structure after 10 years from the construction. in the example the total chloride content was evaluated at depth intervals, ∆c, of 5 mm. from the concrete cover depth measurements, carried out in situ, a frequency analysis can be performed (the depth intervals should be equal to those of the chloride profiles) (figure 2). � � � � � � �� �� �� �� �� �� )u ht xh qf \� �� � &rqfuhwh�fryhu�wklfnqhvv��pp� figure 2. imaginary frequency analysis (grey symbols) and probability density function (black line) of the concrete cover thickness. in order to evaluate the extent of the depassivated bars, initially the variability of the concrete cover 345 federica lollini acta polytechnica ctu proceedings profile ccl (% mass of cement) pcorr, i 1 0.44 0.14 2 0.98 0.99 3 1.33 1 4 1.41 1 5 1.44 1 table 1. chloride contents at the depth c = 22.5 mm and calculated values of pcorr, i. ∆c [mm] 0 − 5 5 − 10 10 − 15 15 − 20 20 − 25 25 − 30 30 − 35 35 − . . . fc [%] 10.2 5.6 6.5 4.6 13 15.7 17.6 26.9 pcorr [%] 100 100 99.9 97 82.5 73.2 55 41 rdep [%] 10.2 5.6 6.5 4.5 10.7 11.5 9.6 3.2 table 2. proportion, fc, of imaginary cover depth measurements falling within the range ∆c, probability of occurrence of corrosion, pcorr, and extent of corroding reinforcement, rdep, in each depth ∆c. thickness, c, is neglected and it is considered that the bars have the same concrete cover (i.e. the concrete cover of all the bars falls within the same cover depth interval, ∆c). moreover, it is initially supposed that, at the bars depth, c, a unique value of the chloride content, ccl, is available. the probability that corrosion initiated, pcorr , corresponds to the probability of occurrence that the variable clth is equal to the considered value of chloride content, ccl. the probability that corrosion occurred is an estimation of the extent of corroding reinforcement. this means that, if at the bars depth a chloride content lower of 0.2 % mass of cement is measured (that corresponds to the lower limit of the beta distribution), pcorr should be null and hence all the bars would be passive, conversely, if the chloride content were higher than 2 % mass of cement (that corresponds to the upper limit of the beta distribution), pcorr would be equal to 100 % and, hence, all the bars would be actively corroding. as an example, a concrete cover thickness equal to 22.5 mm (as indicated by the vertical dotted line in figure 1b) and the first profile, i.e. profile 1, can be considered; at this depth a chloride content of about 0.44 % mass of cement can be evaluated (horizontal dotted line in figure 1b). at this chloride content corresponds a probability pcorr (= rdep) equal to 14 %, given by the intersection between the horizontal dotted line and the clth cumulative distribution function, as shown in figure 1a. considering that more measurements of chloride content at the bars depth, c, are available (as in the example of figure 1b), pcorr can be evaluated through the equation: pcorr = n! i=1 pcorr, i n (1) where n is the number of chloride profiles and i represents the i−th profile. table 1 reports the chloride contents at the concrete cover c = 22.5 mm, shown in figure 1b, and pcorr, i evaluated for each profile. applying equation 1 it can be evaluated that, for this depth interval, pcorr is about 83 %. however, in a real structure, it has to be taken into account also the variability of the concrete cover thickness, as shown for instance in figure 2 and reported in table 2 in terms of percentage of measurements, fc, falling within different depth intervals. hence, pcorr has to be calculated, through equation (1), for all the concrete cover depth intervals, ∆c, as shown in table 2. rdep, j can be, then, evaluated, for each ∆c, as: rdep, j = pcorr, j · fc (2) where fc is the proportion of cover depth measurements falling within the range ∆c and j is equal to the given interval depth, ∆c. hence, the extent of corroding reinforcement can be calculated by summing up corrosion proportions over all the depth intervals: rdep = m! j=1 " #fc, j · $ n! i=1 pcorr, i n % j & ' (3) where n is the number of chloride profiles and i represents the i−th profile, m is the number of cover depth intervals, ∆c, and j represents the j−th depth. by applying equation 3 to the given example, a value of rdep about 62 % can be evaluated. 2.2. probabilistic approach to evaluate the extent of corroding reinforcement, a probabilistic approach, similar to that used for the design of new rc structures, can be also used. in the probabilistic approach for the design of new structures, the extent of corroding reinforcement corresponds to the probability, pdep, that the limit state 346 vol. 33/2022 corrosion damage for rc structures profile 1 2 3 4 5 dapp [10 − 12 m2/s] 0.4 0.7 1 1.2 1.6 cs [% mass of cement] 2.8 3.45 3.6 3.4 3 table 3. diffusion coefficient, dapp, and surface chloride concentration, cs, of imaginary chloride profiles. parameter unit type of distribution value m σ dapp [10 − 12 m2/s] normal 0.98 0.46 cs [% mass of cement] normal 3.25 0.33 c [mm] normal 27.9 14.3 clth [% mass of cement] beta 0.6 0.15 t [year] 10 table 4. values of the parameters and types of statistical distribution used as inputs in the probabilistic approach (m = mean value, σ = standard deviation). equation for the initiation of corrosion, g, is not satisfied, i.e. that pdep{g} < 0. for chloride-induced corrosion, the limit state equation for the initiation of corrosion can be described by means the solution of fick’s second law: pdep = {clth − ccl (x, t)} = p ( clth − cs ) 1 − erf c 2 * dapp t +, < 0 (4) where: clth is the critical chloride threshold, cs is the surface chloride concentration, c is the concrete cover depth, dapp is the diffusion coefficient and t is the time. the variables clth, cs, c and dapp are stochastic variables, whilst the time, t, is a deterministic variable. to evaluate the actual corrosion conditions of an existing structure, reliable values for the input parameters need to be provided. as far as the time, t, is concerned, it corresponds to the time between the year of construction and the considered time (in the example, 10 years). the critical chloride threshold, as previously observed, can be described for carbon steel bars through a beta distribution, with a mean value of 0.6 % mass of cement. the other parameters are related to the considered existing structure and to its exposure conditions and can be determined, if results of a detailed inspection are available, from the experimental measurements. the concrete cover distribution can be evaluated by interpolating the experimental measurements and determining the distribution parameters, i.e. the mean value and the standard deviation. from the imaginary concrete cover thickness shown in figure 2, a normal distribution, with a mean value of 27.9 mm and a standard deviation of 14.3 mm can be determined (black line in figure 2). the interpolation of the chloride content at different depths, at time, t, by means of the second fick’s law, allows to calculate, for each profile, a value of dapp and cs. from the chloride profiles shown in figure 1b, data summarised in table 3 can be evaluated. from these values, both for the diffusion coefficient and the surface chloride concentration, the probability density function, with its characterizing parameters, can be determined. assuming for both parameters a normal distribution, for the diffusion coefficient a mean value and a standard deviation respectively equal to 0.98 and 0.46 × 10−12 m2/s can be calculated, whilst for the surface chloride concentration a mean value and a standard deviation equal to 3.25 % and 0.33 % mass of cement can be determined. table 4 summarizes the values of the parameters involved in equation 4, to be applied for the evaluation of the extent of corroding reinforcement of the example. solving equation 4 through a probabilistic technique, as the montecarlo method, with the values presented in table 4, a pdep equal to 61 % can be evaluated. 3. discussion both mathematical approaches allow to evaluate the extent of corroding reinforcement i.e. the percentage of reinforcement likely to be non-passive, for a reinforced concrete structure subjected to chlorideinduced corrosion, by comparing the chloride content at the depth of the reinforcement, measured during an inspection, with the critical chloride threshold, taking into account the variability of each involved parameter. the extent of corroding reinforcement could be evaluated both at the time of the inspection and in the future, by estimating the evolution in time of the chloride content at different depths. the approaches provide only indications on the amount of actively corroding bars, but do not give any indication on their localization. furthermore they might overestimate the percentage of bars where initiation of corrosion really occurred. because of the mechanism of chloride-induced corrosion, in fact, in the surrounding zones of corroding bars that usually have a chloride content higher than the chloride threshold, 347 federica lollini acta polytechnica ctu proceedings the steel benefits of the formation of macrocell that provides cathodic polarization, preventing the initiation of corrosion [3]. the statistical approach appears to be easily applicable since the punctual value of the chloride content is directly compared with the critical chloride threshold, whilst the probabilistic approach requires the elaboration of the data obtained through experimental measurements, i.e. the concrete cover thickness and the chloride profile, to determine the probability density function of the input parameters. however, in principle, the probabilistic method could be used for a preliminary estimation of the extent of corroding reinforcement even in absence of any experimental data derived from an inspection. indeed, the input parameters involved in equation 4 could be determined from the design data and the environmental exposure conditions. as concrete cover thickness, the design concrete cover could be considered, whilst from the information related to the construction materials, i.e. the w/b ratio and the type of cement, and the exposure condition, the diffusion coefficient and the surface chloride concentration could be estimated. of course, in this case, the reliability of outcomes will depend on the uncertainty in the parameters chosen for the calculations. from their application, values of 62 and 61 % of extent of corroding reinforcement were respectively obtained with the statistical approach and the probabilistic approach, suggesting that the two approaches are comparable. the accuracy and reliability of the outcomes of both approaches depends on the available number of experimental measurements of concrete cover thickness and chloride content at different depths and their representativeness of the whole structure. the reliability of these methods could be validated by comparing their results with a careful corrosion potential mapping that allows to detect the depassivated area. the availability of a methodology that allows to evaluate the percentage of corroding reinforcement, provided that it is representative of the corrosion conditions of the whole structure, could help the designer in the choice of the most suitable repair technique. the percentage of depassivated area and its evolution in time will represent the amount of concrete, even that sound, that must be removed in the conventional repair, since the concrete must be removed in all areas where the chloride threshold is reached at the depth of the reinforcement or it is expected to be reached during the design life of the repair. this repair technique, that also requires the careful cleaning of the surface of the reinforcement to remove all chloride contaminated rust, appears to be suitable when the amount of concrete to be removed and replaced is limited. conversely when large amount of sound concrete should be removed, i.e. when the damage is extensive, other repair methods, such as cathodic protection or electrochemical chloride removal, that require the removal of the only cracked concrete, may be considered. finally such kind of approaches can be also useful in validating the outcome of the recently proposed probabilistic models for service life design as, for instance, the fib model code for service life design. in these models the probability of failure, i.e. the probability of occurrence of corrosion, is evaluated as the probability that the initiation limit state function reaches negative values and, in case of chlorideinduced corrosion, it is defined as the difference between the critical chloride threshold and the chloride content at the depth of the bars. the probability of failure has the same meaning of the percentage of corroding reinforcement. hence, the application of the models for service life design to existing structures and the comparison of the results to the real corrosion conditions of these structures might allow a preliminary understanding of their reliability. usually in these models the concrete performances in relation to the resistance to chloride penetration are evaluated through accelerated tests. the results of accelerated tests are then modified through a series of corrective factors that are directly provided by the models themselves. these accelerated tests were usually not carried out for structures built in the past and, hence, to apply these models to an existing structure, attention should be paid in their definition. an inaccurate estimation of the parameters chosen for the calculation could compromise the comparison with inspection results. 4. conclusive remarks this paper presented two approaches to evaluate the extent of corroding reinforcement i.e. the percentage of reinforcement likely to be non-passive, of a reinforced concrete structure exposed to chloride-induced corrosion, from the results of an inspection, provided that its results are representative of the whole structure. in the statistical approach the extent of corroding reinforcement is calculated by summing up corrosion proportions that correspond to the probability that the critical chloride content is equal to the chloride content at a certain depth interval, over all the considered depth intervals. in the probabilistic approach the extent of corroding reinforcement, corresponds to the probability that the difference between the critical chloride threshold and the chloride content at the depth of the bars, both considered as stochastic parameters, reaches negative values. their application to a fictitious example showed similar results, suggesting that the two approaches are comparable. these approaches can be useful to properly plan a restoration intervention, since the outcomes represent the amount of concrete, even sound, that must be removed in the conventional repair, providing then indications on the suitability of such kind of repair technique rather than other treatments, such as the 348 vol. 33/2022 corrosion damage for rc structures cathodic protection or the electrochemical removal of chlorides, that require the removal of the only cracked concrete. finally, these approaches can be used to assess the reliability of the recent proposed models for service life design, by comparing the outcomes of their application to existing structures with the results to the real corrosion condition of these structures. references [1] r. b. polder, w.h.a. peelen, w.m.g. courage. nontraditional assessment and maintenance methods for aging concrete structures technical and non-technical issues. materials and corrosion 63:1147-1153, 2012. https://doi.org/10.1002/maco.201206725. [2] e. schlangen. eco-efficient repair and rehabilitation of concrete infrastructures, cambridge: woodhead publishing, 2018. [3] l. bertolini, b. elsener, p. pedeferri, et al. corrosion of steel in concrete: prevention, diagnosis, repair. 2nd edition. weinheim: wiley-vch, 2013. [4] m. a. el-reedy. steel-reinforced concrete structures: assessment and repair of corrosion. crc press taylor & francis group, 2007. [5] e. redaelli, m. carsana, m. gastaldi, et al. electrochemical techniques for the repair of reinforced concrete suffering carbonation-induced corrosion. corrosion reviews 29(5-6), 2011. https://doi.org/10.1515/corrrev.2011.008. [6] l. bertolini, m. carsana, m. gastaldi, et al. corrosion assessment and restoration strategies of reinforced concrete buildings of the cultural heritage. materials and corrosion 62(2):146-54, 2011. https://doi.org/10.1002/maco.201005773. [7] m. g. stewart. workmanship and its influence on probabilistic models of concrete compressive strength. materials journal 9(4):361-372, 1995. [8] s. lim, m. akiyama, d. m. frangopol, et al. experimental investigation of the spatial variability of the steel weight loss and corrosion cracking of reinforced concrete members: novel x-ray and digital image processing techniques. structure and infrastructure engineering 13(1):118-34, 2016. https://doi.org/10.1080/15732479.2016.1198397. [9] u. m. angst. challenges and opportunities in corrosion of steel in concrete. materials and structures 51(1), 2018. https://doi.org/10.1617/s11527-017-1131-6. [10] j. mattila. corrosion of steel in reinforced concrete structures, final report, european communities, luxembourg, p. 207-209, 2003. [11] c. code. model code for service life design. lausanne: federation internationale du beton, fib. bulletin (34), 2006. 349 https://doi.org/10.1002/maco.201206725 https://doi.org/10.1515/corrrev.2011.008 https://doi.org/10.1002/maco.201005773 https://doi.org/10.1080/15732479.2016.1198397 https://doi.org/10.1617/s11527-017-1131-6 acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0133 acta polytechnica ctu proceedings 26:133–138, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app comparing mori-tanaka method and first-order homogenization scheme in the viscoelastic modeling of unidirectional fibrous composites soňa valentová, michal šejnoha∗, jan vorel czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: sejnom@fsv.cvut.cz abstract. a comparative study of the viscous response of polymer matrix based fibrous composites predicted by the mori-tanaka method and finite element simulations based on the 1st order homogenization theory is presented. aligned basalt and carbon fibers embedded into a polymeric matrix are considered to represent a quasi isotropic and transversely isotropic two-phase systems. while differences in the prediction of the macroscopic elastic response are attributed merely to the properties of the fiber phase, the viscoelastic behavior is largely affected by the selected homogenization method. a stiffer response predicted by the mori-tanaka method for both creep and relaxation tests is observed for both material systems and supports similar finding found in the literature. thus suitable modifications of the original formulation of such two-point averaging schemes are needed for them to be applicable in the multi-scale modeling of generally anisotropic yarns in plane weave textile composites. keywords: viscoelasticity, homogenization, mori-tanaka method, periodic unit cell, transformation field analysis, fibrous composites. 1. introduction a plane weave textile composites made of woven yarns reinforced by unidirectional fibers bonded to a polymeric matrix are used in many engineering applications. since often exposed to rate dependent or long lasting variable loading the viscous behavior of the matrix phase should be properly represented to arrive at reliable predictions of generally time dependent macroscopic response of such composites. it has been shown in [1] that random nature of the yarn microstructure may play a significant role in the prediction of its macroscopic viscoelastic response particularly in case of high contrast in material properties of the fiber and matrix phase. therein, the concept of statistically equivalent periodic unit cell (sepuc) [2, 3] was exploited to address adequacy of the often assumed periodic hexagonal arrangement (pha) of fibers. this issue is also briefly addressed herein. the point is that if sufficiently accurate, the pha model may be used to validate the applicability of the mori-tanaka method because of their correspondence in the elastic regime. following the footsteps of [4] the virtual creep or relaxation tests carried out numerically at the level of yarn could be used to calibrate a suitable macroscopic viscoelastic model to represent the yarn behavior at the level of textile. however, the macroscopically isotropic response of asphalt mixtures assumed in [4] is hardly acceptable for macroscopically orthotropic fibrous composites. this opens the way to fully coupled multi-scale analysis. most often the fe2 computational scheme [5] is used, where the sepuc model at the level of yarns is loaded by increments of the local strain averages derived at the level of yarns for a given increment of the macroscopic load. the sepuc analysis provides the instantaneous homogenized yarn stiffness and corresponding stress increment. such a format of stress update may often be very computationally expensive. a potential route appears in replacing detailed finite element simulations by a suitable micromechanical model such as the mori-tanaka averaging scheme [6] assuming a piece-wise uniform distribution of stresses and strains in individual material phases thus utilizing the so called two-point averaging scheme in case of two-phase materials, see e.g. [7, 8]. however, using the mori-tanaka method may raise some doubts because of the method limitations in connection to time dependent or nonlinear analyses. in particular, much stiffer response delivered by the mori-tanaka method in comparison to unit cell finite element simulations has often been reported. the purpose of this study is to contribute to this subject in the light of purely viscoelastic behavior of the matrix phase and if needed to offer suitable modifications to the original formulation of such micromechanical models while still keeping their computational efficiency. the paper is organized as follows. section 2 introduces the two material systems examined herein. a formulation of the adopted viscoelastic model is presented next in section 3 followed in section 4 by short theoretical description of the two homogenization approaches. the principal achievements are summarized in sections 5 and 6. 133 https://doi.org/10.14311/app.2020.26.0133 https://ojs.cvut.cz/ojs/index.php/app s. valentová, m. šejnoha, j. vorel acta polytechnica ctu proceedings ea et ga gt νa a b n cf [gpa] [gpa] [gpa] [gpa] [-] [gpa]−1 [gpa]−1 [-] [-] carbon fiber 294 13 12 5 0.24 0.67 basalt fiber 69.68 64.82 28.10 26.14 0.40 0.69 epoxy matrix 2.03 2.03 0.725 0.725 0.40 0.0474 0.00214 0.3526 table 1. material properties of individual phases. 2. materials two material systems are examined in the present study to address both the influence of microstructural details and material properties of individual phases. the same epoxy resin was considered for both carbon and basalt fiber composite with their microstructure displayed in fig. 1(a) and (b), respectively. (a) (b) figure 1. example images of local microstructure: a) carbon fibers, b) basalt fibers. reprinted from [9] with permission from begell house publishers. the corresponding material data taken from [9– 11] are listed in table 1. while basalt fibers are quasi-isotropic the carbon fiber composite represents a highly anisotropic system. on the contrary, the contrast in material properties of the fiber and matrix in the transverse plane is relatively small in case of carbon fibers as oppose to basalt fibers. this particular mismatch in material properties will be reflected in section 4.2 when constructing the computational model. 3. local constitutive equations for the selected epoxy resin the last three parameters in table 1 are the model parameters of a simple power law like formula for the relaxation rate1 r(t,τ) = 1 a + b(t− τ)n , (1) where r(t,τ) is the relaxation function representing the stress at time t due to a unit strain applied at time τ and held constant. to facilitate the numerical solution this function is typically approximated by the dirichlet series in the form r(t,τ) = m∑ µ=1 eµ(τ) exp [ τ θµ − t θµ ] , (2) 1note that these specific values of parameters a,b,n require the time t − τ be given in minutes to yield the viscoelastic modulus in [mpa]. where θµ are the a priory selected relaxation times. functions eµ are usually obtained by fitting the relaxation function, e.g. eq. (1), via eq. (2) using the method of least squares. (a) (b) figure 2. a) maxwell chain unit, b) relaxation function. assuming for example the generalized maxwell chain model in fig. 2(a) composed of ten maxwell elements, we get the spring moduli eµ listed in table 2. note that sum of all spring moduli provides the elastic modulus of the epoxy matrix in table 1. for illustration, fig. 2 plots the relaxation functions provided by eqns. (1) and (2). µ θµ [mpa·s] eµ [mpa] 1 2.949089 3.548983×101 2 9.506234 9.965608×101 3 8.905084×101 2.133588×102 4 7.939963×102 3.599087×102 5 6.874448×103 4.325570×102 6 5.709058×104 3.954614×102 7 5.221879×105 2.450281×102 8 4.561310×106 1.384739×102 9 4.654514×107 5.785885×101 10 4.749622×108 5.854287×101 table 2. parameters of maxwell chain model. for the generalized maxwell chain model it holds σ = m∑ µ=1 σµ, (3) where σµ are stresses in individual maxwell elements. integrating the differential form of constitutive equation for each maxwell element over the time increment 134 vol. 26/2020 viscoelastic modeling of unidirectional fibrous composites ∆t and using eq. (3) gives the increment of the total stress ∆σ as ∆σ = ê (∆ε− ∆µ) , (4) where the instantaneous young’s modulus ê and the increment of viscoelastic strain ∆µ read [12] ê = n∑ µ=1 ∆teµ θµ ( 1 − exp [ − ∆t θµ ]) , (5) ∆µ = 1 ê n∑ µ=1 ( 1 − exp [ − ∆t θµ ]) σµ(ti−1), (6) where σµ(ti−1) represents the hidden stress at the beginning of the new time increment. the stress σµ(ti) at the end of the time increment is then provided by σµ(ti) = σµ(ti−1) exp [ − ∆t θµ ] + (7) + eµ ( 1 − exp [ − ∆t θµ ]) (∆ε− ∆µ). extension of the above set of one-dimensional equations to a general three-dimensional space is provided, e.g. in [1]. 4. homogenization schemes in this section we briefly review the mori-tanaka micromechanical model for a two phase material system in the framework of transformation field analysis [13] and the first-order homogenization scheme grounding on the existence of periodic fields. 4.1. mori-tanaka method consider a two-phase composite made of elastic transversely isotropic fibers of a circular cross-section embedded into a viscoelastic isotropic matrix. assuming piece-wise constant local fields in individual phases r = f,m2 allows us to write the local constitutive equations in the form ∆σf = lf ∆εf, ∆σm = l̂m(∆εm − ∆µm), (8) where l is the material stiffness matrix and l̂ represents the dependence on the viscoelastic modulus ê. suppose that the composite is loaded on its outer boundary either by the prescribed displacements (strain control) or tractions (stress control) compatible with macroscopic uniform strain ∆e or stress ∆σ increments. the local strains or stresses in terms of their applied macroscopic counterparts follow from dvorak’s transformation field analysis [13] as ∆εf = âf ∆e + d̂fm∆µm, (9) ∆εm = âm∆e + d̂mm∆µm, (10) ∆σf = b̂f ∆σ − f̂fml̂m∆µm, (11) ∆σm = b̂m∆σ − f̂mml̂m∆µm, (12) 2subscripts f,m stand for the fiber and matrix phase, respectively. where âr,b̂r and d̂rm,f̂rm,r = f,m, are the mechanical strain and stress localization factors and strain and stress transformation influence functions, respectively. the mori-tanaka method gives the mechanical localization factors in the form âm = [ cmi + cf t̂f ]−1 , âf = t̂f âm, (13) b̂m = [ cmi + cf ŵf ]−1 , b̂f = ŵf b̂m,(14) where the so called partial strain and stress localization factors t̂f, ŵf can be expressed in terms of the eshelby tensor ŝ as t̂f = [ i − p̂ ( l̂m − l̂f )]−1 , (15) ŵf = [ i − q̂ ( m̂m − m̂f )]−1 , (16) p̂ = ŝm̂m, p̂ = −l̂m ( ŝ − i ) , (17) where m = l−1 is the material compliance matrix. note that in case of the mori-tanaka method the eshelby tensor ŝ depends on instantaneous properties of the matrix phase. for a cylindrical fiber3 its particular form is available, e.g. in [8]. for a two-phase composite the transformation influence functions are readily provided in terms of the localization factors as d̂fm = ( i − âf )( l̂m − lf )−1 l̂m, (18) d̂mm = ( i − âm )( l̂m − lf )−1 l̂m, (19) f̂fm = ( i − b̂f )( m̂m − mf )−1 m̂m, (20) f̂mm = ( i − b̂m )( m̂m − mf )−1 m̂m. (21) 4.2. 1st order homogenization assuming again the composite be loaded either by the macroscopic uniform strain ∆e or stress ∆σ admits the following decomposition ∆u(x) = ∆e x + ∆u∗(x), (22) ∆�(x) = ∆e + ∆�∗(x), (23) where ∆u∗(x) is the fluctuation part of the local displacement field ∆u(x). the former one is considered periodic to give the volume average of the fluctuation strain ∆�∗(x) equal to zero. the stepping stone in the derivation of fluctuation displacements ∆u∗(x) is the hill lemma given by〈 δ�t∆σ 〉 = δet∆σ. (24) 3note that for cylindrical fibers the eshelby tensor depends on the matrix phase poisson ratio only, which is assumed constant in our study, i.e. ŝ = s . also note that keeping the time increment ∆t constant yields the viscoelastic modulus êm also constant. thus the localization and transformation matrices need to be evaluated only once. 135 s. valentová, m. šejnoha, j. vorel acta polytechnica ctu proceedings where 〈·〉 stands for volume averaging. in the framework of finite element discretization we approximate ∆u∗(x) in terms of their nodal values ∆r to write ∆u∗(x) = n(x)∆r, ∆�∗(x) = b(x)∆r. (25) substituting from eq. (25)2 into (24) yields the final system of algebraic equations in the form [ k11 k12 k21 k22 ]{ ∆e ∆r } = { ∆σ + ∆f 0 ∆f0 } . (26) individual matrices listed in eq. (26) are written as k11 = 1 ω ∫ ω l̂(x) dω, k12 = kt21 = 1 ω ∫ ω l̂(x)b(x) dω, (27) k22 = 1 ω ∫ ω bt(x)l̂(x)b(x) dω, and components of the right-hand side vector are ∆f 0 = 1 ω ∫ ω l̂(x)µ(x) dω, (28) ∆f0 = 1 ω ∫ ω bt(x)l̂(x)∆µ(x) dω. recall that the system of equations 26 is adopted when loading the composite by the prescribed macroscopic stress such as in case of creep. simulating a relaxation test calls for the macroscopic strain ∆e be prescribed. the system of equations 26 then simplifies to (k22)∆r = −(k21)∆e + ∆f0. (29) the increments of macroscopic strains ∆e (σ prescribed) or stress ∆σ (e prescribed) follow from the volume averaging of their local counterparts. further details can be found in [1, 8]. 5. results the first study aimed at comparing the results provided by various unit cell models. figures 3(b-c) show 5-fiber and 10-fiber periodic unit cells (puc). note that these were constructed by simply changing the diameter and slightly also the position of sepucs derived originally in [8, chapter 3] for a carbon-carbon composite to meet the required volume fractions. such puc thus should not be termed sepuc for the present systems. inadequacy of such a simple approach is confirmed by the results plotted in fig. 5 for the basalt fiber composite subjected to both creep and relaxation loading conditions displayed in fig. 4. a highly anisotropic behavior predicted by 5-fiber and 10-fiber puc in comparison to pha model in fig. 3(a) is evident in figs. 5(a,c) showing a different (a) (b) (c) figure 3. periodic unit cells: a) pha model, b) basalt 5 fiber uc, c) basalt 10 fiber uc. time [s] σ (σ )xy = 100 [mpa]xx σ time [s] e xy xxe (e ) = 0.375 [%] (a) (b) figure 4. applied loading: a) creep test, b) relaxation test. response of the two systems when loading the composite along x and y axes, but also in figs. 5(b,d) showing an evolution of a non-negligible shear strain and stress for the creep and relaxation loading conditions, respectively, associated with the two loading cases. while, as expected, the response predicted by the pha model confirms a material isotropy in the transverse plane. further study comparing the predictions provided by the fem simulations and the mori-tanaka thus adopted the pha model only. the results appear in fig. 6. figure 6(a) shows the evolution of the macroscopic normal strain exx caused by the macroscopic normal stress σxx. for both material systems the initial elastic response predicted by the mori-tanaka method appears more compliant in comparison to fem simulations, while the subsequent time dependent evolution of strains confirms much stiffer behavior suggested by the mori-tanaka method. this is supported by the results derived for the shear loading, figure 6(b), as well as for the relaxation tests, figs. 6(c,d) the results clearly show that the predictions depend not only on the material system but also on the applied type of loading conditions. while the differences in the elastic behavior can be associated with the material properties of the fiber phase, the evolution of local creep strain is driven by the selected computational model. a non-uniform distribution of creep strain in the matrix phase caused by the microstructural details, 136 vol. 26/2020 viscoelastic modeling of unidirectional fibrous composites 0 60 120 180 240 300 360 time [s] 0,50 0,75 1,00 1,25 e [ % ] pha 5 fiber uc 10 fiber uc σ xx applied, e xx plotted σ yy applied, e yy plotted (a) 0 60 120 180 240 300 360 time [s] -0,10 -0,08 -0,06 -0,04 -0,02 0,00 e x y [ % ] pha 5 fiber uc 10 fiber uc σ xx applied σ yy applied (b) 0 60 120 180 240 300 360 time [s] 40 50 60 70 80 σ [ m p a ] pha 5 fiber uc 10 fiber uc e xx applied, σ xx plotted e yy applied, σ yy plotted (c) 0 60 120 180 240 300 360 time [s] 0,0 0,5 1,0 1,5 2,0 2,5 3,0 σ x y [ m p a ] pha 5 fiber uc 10 fiber uc e xx applied e yy applied (d) . figure 5. test of sepuc applicability: (a-b) creep test time variation of macroscopic strains due to prescribed macroscopic stresses σxx, σyy, (c-d) relaxation test time variation of macroscopic stresses due to prescribed macroscopic strains exx, eyy 0 60 120 180 240 300 360 time [s] 0,8 1,0 1,2 1,4 1,6 1,8 e x x [ % ] fem pha mori-tanaka basalt fiber carbon fiber (a) 0 60 120 180 240 300 360 time [s] 2,0 3,0 4,0 5,0 6,0 e x y [ % ] fem pha mori-tanaka carbon fiber basalt fiber (b) 0 60 120 180 240 300 360 time [s] 25 30 35 40 45 50 55 60 σ x x [ m p a ] fem pha mori-tanakabasalt fiber carbon fiber (c) 0 60 120 180 240 300 360 time [s] 6 8 10 12 14 σ x y [ m p a ] fem pha mori-tanaka basalt fiber carbon fiber (d) figure 6. comparison of macroscopic response predicted by finite element simulations and mori-tanaka method for both basalt carbon fiber composites: (a-b) creep test, (c-d) relaxation test a) σxx prescribed, b) σxy prescribed, c) exx prescribed, d) exy prescribed. 137 s. valentová, m. šejnoha, j. vorel acta polytechnica ctu proceedings recall fig. 5, and the mismatch in material properties of the fiber and the matrix phase then renders a significantly more compliant macroscopic response in comparison to the mori-tanaka method, which builds on the piecewise constant stress and strain fields in individual phases, recall eqns.(9) (12). 6. conclusions the present paper examined a potential application of the mori-tanaka method in prediction of effective viscoelastic response of polymer matrix based fibrous composite. the results presented in fig. 6 confirmed a relatively stiff response delivered by the mori-tanaka method when compared to finite element simulations already observed in some previous studies. thus for the mori-tanaka method presented in the framework of dvorak’s transformation field analysis to be applicable as a very efficient substitute of computationally demanding finite element simulations in multi-scale analysis requires some modifications to its original formulation. with regard to generally nonlinear material behavior the literature offers a number of potential approaches. reduction of stresses in the reinforcing phase, typically assumed linearly elastic, by introducing an artificial damage parameter is one particular example [7, 14]. slightly different approach has been proposed in [15] for the analysis of quasi-brittle materials where the effect of strain localization was addressed with the aid of interfacial coating to allow for gradual debonding and thus letting the matrix to take all the stresses. another approach offers more refined transformation field analysis where the concentration and transformation matrices are derived from finite element method using rather crude meshes where each element represents a single material [16]. all these approaches will be examined in our future studies depending on the results provided by an accompanied experimental program. acknowledgements the financial support provided by the gačr grant no. 1915666s and by the czech technical university in prague within sgs project with the application registered under the no. sgs19/032/ohk1/1t/11 is gratefully acknowledged. references [1] m. šejnoha, j. zeman. overall viscoelastic response of random fibrous composites with statistically quasi uniform distribution of reinforcements. computer methods in applied mechanics and engineering 191(44):5027–5044, 2002. doi:10.1016/s0045-7825(02)00433-4. 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[11] s. valentová, v. hrbek, j. vorel, m. šejnoha. strength of composite yarn under biaxial loading. acta polytechnica ctu proceedings 15:131–136, 2018. doi:10.14311/app.2018.15.0131. [12] z. bittnar, j. šejnoha. numerical methods in structural engineering. asce press, 1996. [13] g. j. dvorak, y. benveniste. on transformation strains and uniform fields in multiphase elastic media. proceedings of the royal society of london series a mathematical, physical and engineering sciences 437(1900):291–310, 1992. doi:10.1098/rspa.1992.0062. [14] m. šejnoha, r. valenta, j. zeman. nonlinear viscoelastic analysis of statistically homogeneous random composites. international journal for multiscale computational engineering 2(4):645–673, 2004. doi:10.1615/intjmultcompeng.v2.i4.80. [15] j. vorel, v. šmilauer, z. bittnar. multiscale simulations of concrete mechanical tests. journal of computational and applied mathematics 236(18):4882–4892, 2012. [16] g. dvorak, y. bahei-el-din, a. wafa. the modeling of inelastic composite materials with the transformation field analysis. modelling and simulation in materials science and engineering 2(3a):571, 1999. doi:10.1088/2f0965-0393.2f2.2f3a.2f011. 138 https://doi.org/10.1016/s0045-7825(02)00433-4 https://doi.org/10.1016/s0022-5096(00)00027-2 https://doi.org/10.1088/0965-0393/15/4/s01 https://doi.org/10.1615/intjmultcompeng.v8.i1.100 https://doi.org/10.1002/nme.541 https://doi.org/10.1016/0167-6636(87)90005-6 https://doi.org/10.1615/intjmultcompeng.2014011020 https://doi.org/10.14311/app.2018.15.0131 https://doi.org/10.1098/rspa.1992.0062 https://doi.org/10.1615/intjmultcompeng.v2.i4.80 https://doi.org/10.1088/2f0965-0393.2f2.2f3a.2f011 acta polytechnica ctu proceedings 26:134–139, 2020 1 introduction 2 materials 3 local constitutive equations 4 homogenization schemes 4.1 mori-tanaka method 4.2 1st order homogenization 5 results 6 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0112 acta polytechnica ctu proceedings 33:112–118, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague limitations of life cycle assessment of green concretes a state-of-the-art review amina dacić∗, oliver fenyvesi department of construction materials and technologies, budapest university of technology and economics, müegyetem rkp. 3, 1111 budapest, hungary ∗ corresponding author: amina.dacic@epito.bme.hu abstract. there have been numerous researches focusing on the alternatives to conventional concrete regarding environmental benefits and trade-offs. for this purpose, nowadays, the life cycle assessment (lca) is widely used. until now, lca results of conventional and green concretes have shown some limitations, and this paper focuses on addressing them. firstly, the researches have been mostly based on a small number of concretes, which can limit the possibility to quantify the range of emissions for different set of assumptions. the second one is regarding the appropriate selection of functional unit where a divergence between different researches is visible. moreover, it has been recognised that the criteria involved in the investigation of transportation might lead to the variation of results. also, inventory allocation has to be assessed in detail to reach adequate distribution of the environmental impact of the concrete during its life cycle phases. finally, co2 uptake of the concretes throughout the life cycle should be considered and applied adequately to have representative results of overall co2 emissions. it is expected that the differences will depend on construction practices among the countries where the researches are executed. also, the results will surely vary between different system definitions when lca is applied. keywords: co2 uptake, environmental impact, functional unit, green concrete, inventory allocation, life cycle assessment. 1. introduction sustainable development has got numerous definitions, parkin recognised around 200. this reflects the complexity of the concept and its interpretation as well as possible ways of achieving it. however, it is for sure that it should include economic, environmental, and social dimensions. [1] on the other hand, it has been acknowledged that the structures become sustainable when at least one whole life cycle is assessed and both environmentally beneficial manufacturing processes and deconstruction concepts incorporate the development of circular economy solutions. these solutions are usually based on recycling and reuse of construction and demolition waste (cdw). [2] although the construction industry is one of the most significant for the economy, it carries the burden of one of the most harmful to the environment. the contribution of it to the gross domestic product (gdp) is about 9%, and it is providing 18 million direct jobs through around 3 million enterprises in the european union (eu). on the other hand, the negative effects of the industry are numerous, including dust and gas emissions, land depletion and deterioration, energy and non-renewable natural resources consumption, noise pollution and cdw generation. [3] the construction industry is accountable for the consumption of 50% of natural raw materials, 40% of the energy produced and the generation of 50% of global waste. [4] in countries of organisation for economic co-operation and development (oecd) and eu buildings contribute to about 30% of total primary energy consumption and more than 30% of greenhouse gas (ghg) emissions. [5] kofoworal and ghewaala investigated a typical concrete office building in thailand based on life cycle energy (lce) analysis and concluded that operation phase contributes 81.3% to the building’s lce profile. while the manufacturing of the materials was responsible for 16.8% where steel accounted for 42% and concrete for 35% of the initial embodied energy, respectively. the construction, demolition and maintenance phases values were equal to 0.6%, 0.4% and 0.8%, respectively. however, it was recognised that recycling of building materials can contribute about 9% to energy savings. [6] consequently, sustainable construction has been a focus of research for many decades. in that sense focusing on designing, building, and occupying more sustainable structures would benefit all three beforementioned dimensions of sustainable development. regarding the environmental aspect of sustainable development, it would lead to improvement in air and water quality, reduction in energy and water consumption, as well as in waste disposal. on the other hand, economic benefits would incorporate a decrease in operating and maintenance costs and an increase in revenue due to sale price or rental. finally, in a social sense, the enhancement 112 https://doi.org/10.14311/app.2022.33.0112 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 green concretes life cycle assessment of the occupants’ comfort and health, reduction in absenteeism, turnover rate and liabilities would be included. [7] in the case of the construction industry, it has been recognised that sustainable development should be based on a decrease of raw materials, energy and amount of cdw being landfilled. [8] there are significant economic, environmental and social impacts of cdw landfilling. it can reduce the value of the land stock in the area, it can result in the contamination of soil and groundwater due to cdw’s composition and finally put public health at risk. [9] recycling has been recognised as a good practice that includes the reduction of landfilled cdw and the consumption of both energy and raw materials. as a result, recycled aggregate (ra) has been investigated and more widely used in the construction industry. [8] however, all possible alternatives to the conventional concrete should be investigated regarding environmental benefits and trade-offs. for this purpose, nowadays extensively used life cycle assessment (lca) is applied. [10] although until now many methods have been developed focusing on the evaluation of the environmental impact (ei), lca is the most widely used mainly because of its effectiveness to calculate the potential effects that a product, process or service has or may have on the environment during its whole life cycle. [3] 2. green concrete by definition, green concrete is concrete which uses waste material as at least one of its components, or its production process does not lead to environmental destruction, or it has high performance and life cycle sustainability. [11] in the eu one-third of total waste generated is cdw, making it the largest waste stream. [12] cdw may be resulting from construction activities, natural hazards, demolition or rehabilitation works. furthermore, the most substantial amount is corresponding to demolition and rehabilitation works. around 40 to 50% of cdw is made of concrete, asphalt and brick. [9] the most widely used building material in the construction industry is concrete. [13] moreover, it is the second most utilized material after water. in 2009 on a global level, it is estimated that between 13 to 21 billion tons of concrete are produced annually. [4] only in eu in 2018 total concrete produced was about 320 million cubic meters. [14] moreover, concrete plants globally consume 1 billion tonnes of water, 1.5 billion tonnes of cement and 10 billion tonnes of aggregate annually. [9] these numbers are expected to rise in the future as the production of concrete is expected to increase due to the rise in population and urbanisation. by the increase in the production of the concrete, consumption of natural aggregate (na) as concrete’s largest constituent raises accordingly as well. in the countries of eu annual production of aggregate is equal to three billion tons. [13] out of this value, about 45% is used in concrete. [3] hence, the challenge of the availability of na’s sources has been present, which resulted in the placing of the taxes on the use of na in the case of many european countries. [13] the recycling rate of cdw in eu is on average 83%. [15] there is a significant difference regarding this value among countries since some of the countries have established this practice for years and others have still a very low rate. [16] worldwide countries have adopted the strategy of management of cdw expressed by "3r" (reduction, recycling and reuse). moreover, the eu proposed that the best practice would be for preventing the production of waste. possible solutions to the reduction of cdw production are improving materials which are energy and environmentally efficient, usage of high-performance materials (resulting in the reduction of the amount of material used) and enhancing the durability of buildings (resulting in the increase of the life span of the raw material used). [2] on the other hand, the percentage corresponding to the cdw compared to the total utilisation of aggregate differs between countries, but there is no case where it can completely replace the use of nas. [16] cdw is recognised as one of the most significant solid waste streams and recycling of it into useful components of concrete is seen as an important part of sustainable development. however, there has been extensive research on the inclusion of different types of industrial waste into green concretes either as an aggregate or a binder. the precondition to the usage of such materials is their environmental acceptability and technical adequacy. [10] recycling of cdw can have benefits in reducing the amounts of it disposed to landfills and the preservation of natural resources. this potential can be enhanced if the ra is not just applied to lower quality products but also as a component of structural concrete. even though the research has been extensive in the area of ra, its usage for higher quality products has not been widely recognised. due to some national practices, ra is used for granular base or sub-base applications, embankment construction and earth construction works. [13] moreover, in the researches made, it was shown that na could be successfully substituted by ras in producing concrete that fulfils the requirements for structural use. on the other hand, the application for structural purposes has been recognised as feasible in the both commercial and technical sense. [17] waste recycling has been largely developed for about three decades now. as a result of this activity, many regulations have been established. as an example, the eu directive defines some conditions that waste must fulfil to be considered as a by-product. these conditions are: • further usage of the object or substance is definite; • the object or a substance is generated as an integral part of a production process; 113 amina dacić, oliver fenyvesi acta polytechnica ctu proceedings • the object or substance is usable excluding any further processing other than usual industrial practice; • further usage is legitimate meaning that it fulfils all necessary environmental, health and product requirements for that specific use and would not cause overall negative environmental or human health effects. [18] moreover, the design and process of recycling should take into account the saving of raw materials and energy and decreasing the potential of environmental pollution. consequently, recycling and reuse of cdw offer an important opportunity for sustainable development in the construction industry. [2] 3. life cycle assessment back in the late 1960s and early 1970s, environmental issues were addressed extensively, such as resource and energy efficiency, pollution, and solid waste. during that period, studies on ei of consumer products also started which were mainly based on simple principles to be able to compare two products. as a result of this soon enough, it was recognised that the biggest ei is not connected to the products’ use but rather to its production, transportation, or disposal. while these studies were evolving, there was a need for standardisation of the assessment tool, which would be widely applied. finally, the international organization for standardization (iso) published iso 14040 (2006) and iso 14044 (2006) standards. based on these standards, the methodology consists of four main steps which are the definition of a goal and scope, the inventory analysis, the impact assessment and the interpretation [3]. lca is a methodology used for evaluation of the environmental burden of the processes, products or services throughout their life cycle. [13] in the case of green concretes, lca can be a useful tool for generating indicators regarding ei in source separation, waste treatment and recycling technologies. however, it was recognised that aggregates’ recycling differs between developing and developed countries in terms of production and transportation methods. moreover, resources availability can vary among countries. for example, in the case of china, the shortage of resources has become a serious issue currently, and the possibility of alternatives to conventional concrete is of significant importance. furthermore, source locations are mainly on a longer distance from urban areas which then puts a greater environmental burden on conventional concrete [8]. on the other hand, many european countries have been employing regulations to answer to the decreasing availability of the na resources such as placing taxes on the use [13]. as well as focusing on recovery of cdw by implementing the waste framework directive 2008/98/ec aiming at 70% recycling rate by 2020 [12]. 3.1. quantity of concretes one of the limitations that have been acknowledged considering the lca of conventional and green concretes is that the researches are based on a small number of concretes. the range varies mainly between 3 to 18 mix designs. [4] moreover, the researchers mainly focused on the certain green solution of concretes, generally considering recycled concrete aggregate substitution to nas. it has been concluded that overall there are five types of green concrete on which the researches have been focused: 1. recycled aggregate concrete (rac) for which as aggregates, ras generated from cdw are used; 2. blended cement concrete in ordinary portland cement (opc) in the binder is replaced with a variety of supplementary cementitious materials; 3. alkali activated concrete where no opc is used but materials which are rich in alumina and/or reactive silica; 4. high performance concrete which is considered as a concrete of improved mechanical and durability properties compared to one generated using opc; 5. bio-concrete which is produced mainly using various bacteria. [19] these five types of concrete can have a beneficial behaviour in ei sense compared to traditional concrete. furthermore, most of the researchers have used the conventional mixture design method in case of rac, which is based on replacing a certain percentage of na with ra based on equivalent volume or mass [20]. in advance, it has been noticed that the amount of cement is increased in rac mixes to achieve those properties of natural aggregate concrete (nac) [21]. eventually, this could lead to more co2 emissions. this is why researchers worked on developing some alternative mix design methods that could lead to improved performance of racs and better and fairer comparison to nac. the alternative to the classical approach of mixture design in the case of the rac, equivalent volume mortar method (emv) has been proposed. this method is based on the fact that rac is two-phase material including na and mortar, so it seeks to use the same amount of total mortar volume in mixtures of both nac and rac [22]. other than mixture design methods, the development of the different mixing methods mainly based on the sequential mixing approach in order to improve properties of rac have been introduced [23]. furthermore, there have been just a few researches regarding the effect of mixing methods on lca of rac, which can lead to possible better performance of green concretes on both material and lca level. 3.2. functional unit another limitation, the functional unit (fu) is used as a key factor influencing lca. fu is a quantified 114 vol. 33/2022 green concretes life cycle assessment performance of a product system for use as a reference unit. [24] as the definition implies, it is an important factor for comparison of lca results for investigating alternative solutions for a certain product. that applies to the production of concrete also. so, many researchers based their lca of green concretes on the functional units (fus). however, the selection of fus differs, including volume of concrete, combining two or more variables such as volume and strength; volume, strength, durability; and strength reliability. the biggest disadvantage of having different fus in the conducted researchers is that these results can not be compared. [25] by many researchers volume of concrete as a fu was found as not reliable since it does not take into concern that for 1 m3 nac and rac compressive strength results are not equivalent. on the other hand, fu per strength, which combines both volume and strength, neglects the durability properties of the concrete. [26] however, it has been recognised that fu has to incorporate the workability, strength and durability/service life differences between investigated concretes. [3] 3.3. transportation non-renewable energy consumption and global warming potential (gwp) as ei in the case of concrete significantly depend on the transportation scenarios. [27] ding, xia and tam showed in their study based on data in china that aside from cement proportion, transportation is the top contributor to the co2 emissions and energy consumption for both conventional and rac. in the conducted sensitivity analysis, it was concluded that longer delivery distances for na lead to the possibility of decreasing ei of rac. so, by increasing the replacement ratio of na by ra, there is a possibility of reducing the negative ei of concrete. [8] on the other hand, marinković et al. based their lca in serbia, and the results showed that the total ei of conventional and rac depend on the na and ra transport distances and types. since the ei of the aggregate and cement production phase gave slightly disadvantageous results for rac compared to conventional concrete in their research, the transportation phase was recognised as a possible source for reducing the ei of rac. [13] knoeri, sanye-mengual and althaus examined rac mixtures according to laws, standards, and construction practices in switzerland. they concluded that the additional transportation distances more than 15 km result in higher ei for rac than conventional concrete. based on the present market mixtures in switzerland, the beneficial results can be expected for rac compared to conventional one if supplementary cement and transportation are kept limited. [28] 3.4. inventory allocation life cycle inventory (lci) corresponds to the collection of data, all inputs and outputs of the considered life cycle phases. [9] in green concretes, the research should adequately address inventory allocation for the recycled and/or reused products. for ra, this allocation is important to be suitably distributed between waste management and material production. [4] in general, for the inventory analysis for lca information can be collected from industries involved usually by questionnaires, publicly available annual environmental reports (ers), and environmental product declarations (epds), lca related journals and lca databases (e.g. ecoinvent, gabi). the most precise data can be obtained from questionaries. however, generally all the data can not be collected this way, so alternative beforementioned sources are also used. [3] when it comes to waste recycling, the main issue of the lca is the allocation process. in the present practice to the waste production, no environmental burdens are assigned, if we disregard waste disposal, because it is not generated on purpose. but when we consider waste as a by-product, it has been recognised that some allocation coefficient has to be applied. however, there is no consensus on the correct allocation procedure in this case in the previous researchers. so, it rather relies on the choice of the researcher, and since it can have an important effect on the results, it is one of the most debatable issues in lca of green concretes. according to the abovementioned iso standards when there is a possibility of several allocation procedures, it is recommended to conduct a sensitivity analysis demonstrating the impact of different choices. [18] 3.5. co2 uptake to evaluate the portion of ghg that concrete is responsible for, it has been acknowledged that it is needed to adequately cover the net balance of ghg in the use of concrete as a construction material. this implies that co2 uptake throughout the whole product life cycle shall also be considered. so, as the emissions are emerging by the concrete production, there is a parallel co2 uptake process going on in the already produced concrete. however, when the uptake is considered, there is one corresponding to primary concrete product (i.e. concrete used in the construction of the structures) and the secondary resulting from demolishing and crushing the primary one. this secondary product can be landfilled, used as a road base material or ra in green concretes. [29] if the co2 uptake is not considered in the lca of concrete, it can lead to the overestimation of the emissions. co2 is absorbed during the carbonation of the concrete during both the primary and secondary product phase. [4] collins proposed the formula for the calculation of co2 uptake from the atmosphere: co2uptake = x · copc · cao · r · a · m (1) where x corresponds to calculated depth of carbonation equation (2), copc is amount of opc in binder and cao refers to calcium oxide content of 115 amina dacić, oliver fenyvesi acta polytechnica ctu proceedings opc. while r is a fraction of cao in fully carbonated opc, a denotes surface area of the exposed concrete and chemical molar proportion of co2/cao is assigned by m . moreover, carbonation depth can be calculated using the following formula: x = k √ t (2) where k is carbonation rate coefficient which covers the chemical and physical characteristics of the concrete and environment (can be determined experimentally or measured on the structure) and t is exposure time. [30] from the above-presented equations, it can be concluded that the co2 uptake is dependent on the properties of the concrete such as cement content, co2 in the atmosphere which relies on the exposure condition and diffusion process between the atmosphere and the concrete effected by the exposed surface. [4] full carbonation of the concrete, hence, can lead to a substantial reduction in the gwp. however, this complete carbonation does not seem like a realistic outcome, since by researchers 75 % has been concluded as a maximum value. [31] moreover, just 11 % of theoretical maximum co2 uptake has been recognised as possible during the primary production phase. [29] however, when concrete is considered as a part of the cdw, demolishing and crushing results in the smaller particle size henceforth larger exposed surface area. [31] further investigation has to be conducted regarding the co2 uptake for green concretes, especially when ra is an ingredient of it. 4. conclusions lca is the standardised tool for determining ei of a product or process, and it has been the most widely used one for this purpose. some limitations have been recognised when it comes to its applicability to the evaluation of concrete’s performance. this is especially the case when alternative solutions to conventional concrete are assessed. although the research in this field has been present for some decades now, there are some parts of the lca of the concrete where consensus between researchers has not been made. consequently, the large-scale assessment of the ei of concretes, especially green concretes is scarce. in further researches, special attention should be paid to incorporating a larger quantity of investigated concretes. it is essential to investigate different mix design methods, especially since they may offer better performance in both material and lca sense for green concretes. on the other hand, one of the most critical questions when concretes’ lca is conducted is fu. the variety of fus used in research leads to the limited possibility of comparing the results of the different studies. this is one of the main issues on which consensus should be made to have a more precise picture about ei of concrete. as expected, transportation takes a significant portion of the gwp of concrete. moreover, after the cement production phase, it is the most influential factor. the results showed that by reducing the transportation distance for ra compared to na beneficial results can be conducted in the case of green concretes. in lca of any product or process, inventory allocation is a very important phase. it can lead to significantly overor underestimated results if not adequately conducted. moreover, in the case of green concretes, special attention has to be paid to the usage of materials as by-products. finally, to have representative results of ghg emissions of the concrete adequate net balance calculation has to be made. in that sense, co2 uptake takes a significant role. it is of great importance to consider both primary and secondary (one emerging after demolishing and crushing of the first one) product’s co2 uptake. co2 is absorbed during the carbonation of concrete as both primary and secondary product. if this net balance is not appropriately calculated, the results lead to overestimating ghg emissions of concrete. to sum up, lca of alternatives to conventional concretes require more detailed evaluation to have representative results. further investigations for sure have to be conducted, which will focus on the abovementioned limitations for improving the lca tool for the evaluation of concrete. references [1] s. parkin. sustainable development: the concept and the practical challenge. proceedings of the institution of civil engineers civil engineering 138(6):3-8, 2000. https://doi.org/10.1680/cien.2000.138.6.3. 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https://doi.org/10.1061/(asce)0899-1561(2009)21:10(601) https://doi.org/10.1016/j.conbuildmat.2020.121490 https://doi.org/10.1016/j.conbuildmat.2019.03.078 https://doi.org/10.1016/j.cemconcomp.2012.01.004 https://doi.org/10.1016/j.jobe.2020.101173 https://doi.org/10.1007/s11367-012-0544-2 https://doi.org/10.1021/es401775w https://doi.org/10.1007/s11367-010-0191-4 amina dacić, oliver fenyvesi acta polytechnica ctu proceedings [31] s. butera, t. h. christensen, t. f. astrup. life cycle assessment of construction and demolition waste management. waste management 44:196-205, 2015. https://doi.org/10.1016/j.wasman.2015.07.011. 118 https://doi.org/10.1016/j.wasman.2015.07.011 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2021.31.0045 acta polytechnica ctu proceedings 31:45–52, 2021 © 2021 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague vehicle/track interaction under the conditions of high speed railway operation jiří šlapák∗, tomáš michálek university of pardubice, faculty of transport engineering, department of transport means and dignostics, studentská 95, 532 10 pardubice, czech republic ∗ corresponding author: jiri.slapak@upce.cz abstract. this paper deals with the issue of dynamic effects of a high-speed railway vehicle on the track. some aspects of vehicle/track interaction at higher speeds are described. in this context, a multi-body model of high-speed railway vehicle was created and several simulation scenarios were performed with this model. furthermore, the simulations were evaluated with a focus on selected characteristics of the vehicle/track interaction. keywords: vehicle/track interaction, dynamic wheel force, damaging effect, multi-body simulation. 1. introduction high speed railway operation makes demands on all structural subsystems (infrastructure, energy, rolling stock, control command and signaling) of the railway system which have to meet relevant requirements of the technical specifications for interoperability (tsis). in the field of vehicle/track interaction, these requirements are connected especially with (vertical) wheel forces and vehicle running stability, i.e. with straining of the track. this paper is focused on the influence of vehicle speed on selected characteristics of the vehicle/track interaction representing damaging effects of the vehicle on the track. by means of multi-body simulations of vehicle running performance, a quantification of these damaging effects can be performed. 2. vehicle/track interaction at higher speeds during the run of a railway vehicle on a real track, track irregularities represent a kinematic excitation of the mechanical system of the vehicle. a direct consequence of this effect is an oscillation of the wheel forces (acting between individual wheels and rails in vertical direction) around their static values during the run. the static values of the wheel forces are given by the relevant axle load; amplitudes of the oscillations are influenced with the unsprung masses, vehicle speed and the track irregularities [1]. in order to minimize the dynamic interaction between the vehicle and the track in vertical direction, the axle load of the highspeed vehicles is limited. according to the relevant tsis, the maximum axle load of the vehicles intended for running at speed 250 km/h and more is 17 t [2]; locomotives with the axle load up to 22.5 t can be used up to 230 km/h. a minimization of dynamic components of the (vertical) wheel force interaction can be reached practically only by the reduction of unsprung masses in the vehicle running gear. another aspect of the vehicle/track interaction � the running stability � is related to the safety as well as running comfort. safeguarding of a stable run of the vehicle under all relevant conditions (equivalent conicity) is a necessary condition for the high speed operation. the technical solutions used to improve running stability (very rigid wheelset guiding, effective damping of bogie yaw motion etc.) usually lead to increased damaging effects of the vehicle run on rails in curves. however, the running stability is not the object of this paper. guiding forces and related frictional work in wheel/rail contact (creeps and tangential creep wheel forces) cause the damaging effects of rails in curve. in addition to the technical solution of the vehicle, the damaging effect also depends on the operating conditions of the vehicle (radii of curves, vehicle speed). 3. multi-body simulations for prediction of railway vehicle behaviour under the specific operation conditions and for the first consideration about vehicle/track interaction, a multi-body simulations of railway vehicles can be used. in this case, the high speed operation is the specific condition. the one of the advantage of simulations is that it can predict vehicles running behaviour if the running tests on a real railway vehicle can’t be performed. the suitable model of railway vehicle must be created for multi-body simulations. also the simulation scenarios (track properties) must be designed. the used model and scenarios of simulations correspond to the high speed railway conditions. the multi-body simulations can be performed in a various commercial software, for example adams/rail, vampire, simpack. of course, it is possible to create own simulation program. for purpose of this paper, the multi-body simulations of railway vehicle running were performed using the simulation software simpack 9.9.2. 45 https://doi.org/10.14311/app.2021.31.0045 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en jiří šlapák, tomáš michálek acta polytechnica ctu proceedings 3.1. multi-body model of railway vehicle for the purpose of this work, a multi-body model of a high-speed railway vehicle was created. this model represents one passenger coach. because the parameters of high-speed vehicles are different than parameters of conventional vehicles operating in the czech republic, the multi-body model represents chinese railway high-speed (crh) vehicle, which is described in [3]. selected fundamental parameters of the multi-body model are listed in table 1. parameter value unit mass of the carbody 35880 kg mass of the bogie frame 3300 kg mass of the wheelset 1751.8 kg bogie distance 17.5 m wheelset distance 2.5 m vertical stiffness of the primary spring (per wheel) 1.176 kn/mm vertical stiffness of the secondary spring (per side of bogie) 0.190 kn/mm table 1. fundamental parameters of the multi-body model of the railway vehicle. the wheel profile ore s1002 and the rail profile 60e1 were used. the rail inclination of 1:20 was considered to achieve better running stability of the vehicle model. for wheel/rail contact modeling, the method fastsim and the friction coefficient value f = 0.3 was selected. the reference track irregularities were used for all multi-body simulations. because of realistic modeling of the track, the model of elastic track foundation was used in order to describe the railway sleepers and elastic joint between a sleeper and solid ground. 3.2. simulation scenarios for the purpose of this paper, four scenarios of simulations were designed. these scenarios represent typical operating conditions on high speed railway lines. in all senarios, the 2200 meters long track sections were created. first 200 meters long part of each track section was considered without irregularities. then the reference track irregularities begin and continue until the end of track section. 3.2.1. influence of running on a straight track in the first scenario, each simulation of railway vehicle running was performed on the same straight track but at a different vehicle speed. specifically, the vehicle speeds are given in table 2 for this scenario. the minimum speed (160 km/h) was chosen in order to perform simulations at a vehicle speed higher than the czech conventional track lines allow. the maximum vehicle speed (350 km/h) was chosen with regard to the assumption of a maximal operation speed of 320 km/h. for vehicle speed higher than 300 km/h, the speed of the test run of railway vehicle must be increased by 30 km/h according to [5]. v [km/h] 160 200 225 250 275 300 320 350 table 2. vehicle speed parameters v of the first simulation scenario. 3.2.2. influence of cant deficiency the second scenario deals with railway vehicle running on curved tracks. the same constant radius of curve was set for all the tracks. for high speed railway vehicle operations, the large radii of curves are typical. the value of curve radius r = 4320 m was selected for these simulations. but the vehicle speed was changing for each simulation of running vehicle, it means that a cant deficiency i was changing too. the cant deficiency i was calculated according to the stated conditions and equation 1. for this scenario, the simulation parameters are listed in table 3, where d is the cant. i = 11, 8 · v 2 r −d (1) d [mm] i [mm] r [m] v [km/h] 180 -110.1 4230 160 180 -81.4 4230 190 180 -70.7 4230 200 180 -41.7 4230 225 180 -9.3 4230 250 180 26.6 4230 275 180 65.8 4230 300 180 99.7 4230 320 180 154.6 4230 350 table 3. the parameters of cant d, cant deficiency i, curve radius r and vehicle speed v of the second simulation scenario. 3.2.3. influence of curve radius another scenario is based on a constant value of cant deficiency i = 100 mm. here the radii of curves were changing. for each case of curve radius r, the vehicle speed v was calculated according to equation 2 and the values of cant d and cant deficiency i were constant. the simulation parameters are listed in table 4 for this scenario. v = √ (d + i) · r 11, 8 (2) 46 vol. 31/2021 vehicle/track interaction d [mm] i [mm] r [m] v [km/h] 180 100 3800 300.3 180 100 4100 311.9 180 100 4400 323.1 180 100 4700 334.0 180 100 5000 344.5 table 4. the parameters of cant d, cant deficiency i, curve radius r and vehicle speed v of the third simulation scenario. 3.2.4. influence of specific conditions around the speed limit of 300 km/h the last scenario of simulations was focused to show one of the specific condition of high speed railway operation according to tsi [2]. in the tsi, there is the table with cant deficiency conditions for high speed vehicle operation. the table says that the maximum value of the cant deficiency is i = 153 mm for vehicle speed v ≤ 300 km/h. however for vehicle speed v > 300 km/h, the maximum value of the cant deficiency i = 100 mm is defined. these conditions apply to railway vehicles and locomotives for passenger transport. so the value of cant deficiency is abruptly changed at this specified vehicle speed. the aim of this scenario is to investigate what happens around this vehicle speed. the radii of curves were defined in order to set the vehicle speed around the value v = 300 km/h. for this scenario, the equation 2 was used to calcuate the vehicle speed and the simulation parameters are listed in table 5. d [mm] i [mm] r [m] v [km/h] 180 153 2800 281.1 180 153 3000 291.0 180 153 3150 298.2 180 100 3800 300.3 180 100 4000 308.1 180 100 4200 315.7 table 5. the parameters of cant d, cant deficiency i, curve radius r and vehicle speed v of the fourth simulation scenario. 4. simulation results the multi-body simulations of running vehicle were performed according to the conditions for each scenario. time records of quantities characterizing the dynamic effects of a running vehicle on the track were exported for each multi-body simulation. then the exported data were processed in matlab software. figure 1 shows the marks of individual wheels of the vehicle model. the letters l and r indicate the left and right sides of the vehicle. figure 1. selected marking of individual wheels of the vehicle model. 4.1. results from straight track the vertical wheel forces q were monitored in all cases of straight track simulations. the mean value of vertical wheel force represents the static vertical wheel force qstat, its value does not depend on vehicle speed. in order to capture the dynamic effect of this force, it is necessary to focus on the oscillation around the mean value, i.e. the static value of the vertical wheel force qstat. the oscillating part of vertical wheel force is marked as ∆q. for the purpose of this paper, the force ∆q is called the partial dynamic wheel force. the dynamic wheel force qdyn is expressed in equation 3. qdyn = qstat + ∆q (3) this means that the recorded vertical wheel forces were the dynamic wheel forces qdyn. the 99.85th percentile of the time record of vertical wheel force was used to evaluate the dynamic wheel force qdyn. this evaluation was used for simulations performed at different vehicle speeds (table 2). the dependence between vehicle speed v and mean dynamic wheel force qdyn_m is shown in figure 2. the mean dynamic wheel force qdyn_m was calculated as mean value of dynamic wheel forces qdyn on all wheels of the vehicle. this calculation was performed for each value of vehicle speed. figure 2. influence of vehicle speed v on dynamic wheel force qdyn under the conditions of straight track. figure 2 shows that the mean dynamic wheel force qdyn_m progressively increases with the vehicle speed v . however, according to the british standard gm/tt0088 [1], the dynamic wheel force is linearly dependent on the vehicle speed. the standard defines a force called p 2, which represents the dynamic wheel force and it is described by the equation 4. p2 = qstat + a ·v (4) 47 jiří šlapák, tomáš michálek acta polytechnica ctu proceedings figure 3. influence of vehicle speed v on partial dynamic wheel force ∆q under the conditions of the straight track. the parameter a in equation 4 depends on the unsprung mass in running gear of the vehicle [1]. the second part of equation 4 (a ·v ) corresponds to the partial dynamic wheel force ∆q, it means that the partial dynamic wheel force is a function of the vehicle speed. the influence of vehicle speed v on partial dynamic wheel force ∆q is shown in figure 3. the linear approximation of the dependence of the partial dynamic wheel force on the vehicle speed shows the coefficient of determination r2 = 0.895. the linear function is shown in figure 3 by blue dashed line. the second degree polynomial (quadratic) approximation gives a better result. there is a coefficient of determination r2 = 0.997. this means that the second degree polynomial is a better function for approximating the dependence of the dynamic wheel force on the vehicle speed for high speed railway operation. ∆q = 0.0008 ·v 2 + 0.01855 ·v (5) equation 5 is the polynomial approximation, which is shown in figure 3 by red dashed line. the parameters of this equation are specific to the simulation conditions. for example, the parameters of equation 5 can be changed by various track irregularities. 4.2. results from curved tracks the situation of vehicle/track interaction in curved track is different. the track is significantly loaded in the lateral direction due to the effect of centrifugal (inertia) force. the track leads the railway vehicle through the curve. in addition to the force interaction, there is a sliding movement in the wheel/rail contact. a special characteristic called wear number wn is used to describe the effects of a railway vehicle on a curved track. in general, the wear number represents a specific friction work and it is used to evaluation of rail wear on a curved track, for example [4]. the wear number is defined by tangential contact forces t and creeps γ in the longitudinal (x) and lateral (y) directions. the formula for calculating the wear number is given by equation 6. wn = |tx ·γx| + |ty ·γy| (6) the mean value of wear number was evaluated. to evaluate the wear number, the only part of the track where the whole vehicle was in the curve was selected. this part of the track starts at 700 m and ends at 1700 m, so only 1000 meters of the track were evaluated. 4.2.1. influence of cant deficiency in the first case of a multi-body simulation of railway vehicle in a curved track, the curve radius r and the value of cant d were constant. the vehicle speed v was increased from 160 km/h to 350 km/h in individual steps, as well as the cant deficiency i (according to equation 1). the simulation conditions are described in section 3.2.2. figure 4. dependence of the sum of the wear number w n values of all vehicle wheels on vehicle speed v at constant track parameters (cant, curve radius). firstly, for a very simplified introduction to this scenario, the dependence of the sum of the wear number wn values of all vehicle wheels on vehicle speed v is shown in figure 4. the figure shows that the wear number value increases progressively with the vehicle speed under conditions of constant curve radius. a very similar dependence is shown in figure 6. this figure shows the effect of the cant deficiency 48 vol. 31/2021 vehicle/track interaction figure 5. dependence of the wear number w n values on all outer wheels of the vehicle on the cant deficiency i at constant track parameters (cant, curve radius). i on the wear number wn. negative values of cant deficiency represent the cant excess. figure 6. dependence of the sum of the wear number w n values of all vehicle wheels on cant deficiency i at constant track parameters (cant, curve radius). according to figure 6, the sum of the wear number wn values increases with increasing cant deficiency i. this applies to the positive values of cant deficiency. but for the cant excess values (negative values of cant deficiency), the wear number values are in a narrow range. the situation where the values of the sum of the wear number were calculated for individual sides of railway vehicles is shown in figure 7. the figure shows the sum of wear number of the four outer wheels (left side in figure 1) and the four inner wheels (right side in figure 1) of the railway vehicle. the chart in the figure shows that the values of the sum of wear number on the inner wheels are almost symmetrical around the value of the cant deficiency i = 0 mm with a polynomial character. it is obvious that the wear number values on the outer wheels are bigger than on the inner wheels. the wear number values of the outer wheels for a positive values of the cant deficiency can be approximated using a quadratic function without constant parameter. because the parameters of a quadratic function may vary depending on the vehicle/track conditions (vehicle construction, track irregularities), the type and shape of the function is important. the quadratic function is shown in figure 7 by a gray dashed line. according to this figure, the sum of the wear number on the outer wheels of the railway vehicle for the cant deficiency i = 110 mm is about twice as large as the sum of the wear number for the cant deficiency i = 0 mm. figure 7. dependence of the sum of the wear number w n values on the inner and outer wheels of the vehicle on the cant deficiency i at constant track parameters (cant, curve radius). according to figure 7, it is obvious that the main part of the wear number (specific friction work) is performed on the outer four wheels. that is the reason why this paper will further deal with the situation on the outer wheels of the railway vehicle, which is shown in figure 5. the outer wheels of the front bogie (wheels 1l and 2l in figure 1) are shown in solid lines and the outer wheels of the back bogie (wheels 3l and 4l in figure 1) are shown in dashed line. as figure 5 shows, for cant deficiency values around i = 0 mm, the wear number values for the second (2l) and fourth (4l) wheels are close to zero. then, as the cant deficiency value increases, the value of the 49 jiří šlapák, tomáš michálek acta polytechnica ctu proceedings wear number also increases with a progressive trend. however, the wear number value is about 37 n on the first wheel and about 22 n on the third wheel if the cant deficiency value is i = 0 mm. for these wheels, the wear number values generally increase with increasing cant deficiency values. the wear number values on the first (1l) and third (3l) wheels increase slower then the values on the second (2l) and fourth (4l) wheels. finally, for the cant deficiency value of 150 mm (vehicle speed 344 km/h), the wear number values for the second and fourth wheels are higher than for the first wheel. this is due to high centrifugal force that pushes the vehicle out of the curve. for normal high speed railway operation (the cant deficiency up to 100 mm), the wear number value on the first outer wheel in the front bogie (1l) is the highest value. this means that the largest specific friction work is performed on this wheel. 4.2.2. influence of curve radius in another case of the simulation scenario, the cant deficiency value is constant i = 100 mm. this cant deficiency value is the maximum possible cant deficiency value for a vehicle speed higher than 300 km/h according to the tsi [2]. this means that the vehicle speed v changes with the curve radius r. the specific simulation conditions are descibed in section 3.2.3. for the first introduction of this scenario, the simple chart is shown in figure 8. dependence of the sum of the wear number wn values of all vehicle wheels on the curve radius r is shown under the constant cant d and cant deficiency i conditions. the radius range of the curve from 3800 m to 5000 m corresponds to the range of vehicle speed from 300.3 km/h to 344.5 km/h. it is obvious that the value of the sum of the wear number increases as the value of the curve radius decreases. this trend is also evident in methodology [4] which is focused on the conventional railway operation, so there is the range of the curve radius from 250 m to 1200 m and the values of wear number are much bigger depending on the smaller radius of the curve. figure 8. dependence of the sum of the wear number w n values of all vehicle wheels on the curve radius r at constant value of cant and cant deficiency. for a better view of this situation, the sum of the wear number values on all outer wheels (left side in figure 1) and inner wheels (left side in figure 1) of the vehicle is shown in figure 9. the sum of the wear number on the inner wheels is approximately equal to 10 n for the whole range of curve radii. the main part of the specific friction work is performed in the contacts of the outer wheels with rail as figure 9 shows. for the curve radius r = 5000 m, the value of the sum of the wear number is 75 n, then the value increases to 95 n for the curve radius r = 3800 m. this means that it is necessary to be interested in the wear number values on the outer wheels. figure 9. dependence of the sum of the wear number w n values on the inner and outer wheels of the vehicle on the curve radius r at constant value of cant and cant deficiency. further analysis of the wear number on each outer wheel (wheels 1l, 2l, 3l, 4l according to figure 1) is shown in figure 10. according to figure 10, the largest specific friction work is performed on the first outer wheel (1l) of the vehicle. the values of the wear number on first outer wheel increase from 25 n up to 39 n for the range of the curve radius from 5000 m to 3800 m. in context to these values, it can be assumed that the head checking rail defect will occur [4]. the wear number values on the third outer wheel (3l) have a similar trend as the first outer wheel, but the values are smaller. the wear number values on the second outer wheels of both bogies (2l, 4l) are very similar and the wear number value increases as the radius of the curve increases. in this scenario, the centrifugal forces have the same value for each simulation. thus, the trends of the wear number values are caused by the values of the angle of attack, which increase with decreasing values of the radius of the curve. the absolute values of the wear number depend on other vehicle parameters and track conditions, but it can be assumed that the trends of the graph curves will be the same. 50 vol. 31/2021 vehicle/track interaction figure 10. dependence of the wear number w n values of all outer wheels of the vehicle on the curve radius r at constant value of cant d and cant deficiency i. 4.2.3. influence of specific conditions around the speed limit of 300 km/h the specific conditions of this scenario are defined in section 3.2.4. this scenario of multi-body simulations shows what happens at the vehicle speed v = 300 km/h. the maximum permissible value of the cant deficiency changes abruptly from 153 mm to 100 mm at this vehicle speed. figure 11. dependence of the sum of the wear number w n values of all wheels of the vehicle on the vehicle speed v with the change of the cant deficiency value from 153 mm to 100 mm. the results of the multi-body simulations of this scenario are shown in figure 11. the figure shows the dependence of the sum of the wear number wn of all wheels on vehicle speed v . for vehicle speeds below 300 km/h, the cant deficiency is 153 mm. for vehicle speeds above 300 km/h, the cant deficiency is 100 mm. the radius of the curve varies according to the values of vehicle speed and cant deficiency. according to figure 11 it is obvious that the wear number values for vehicle speed below 300 km/h are greater than the values for vehicle speed above 300 km/h under the conditions of the maximum value of cant deficiency. for the high speed railway operation with a maximum value of cant deficiency, a vehicle speed greater than 300 km/h is more advantageous in terms of specific friction work and track damage. however, larger radii of curve are required. 5. conclusions after comparing some methodologies for the vehicle/track interaction evaluation for conventional railway operation with the vehicle/track interaction under the conditions of high speed railway operation, it’s obvious that the conditions and parameters of high speed railway operation are different from conventional operation. when a high-speed vehicle runs on a straight track, the damaging effect increases as a result of an increase of the dynamic wheel force, which progressively depends on the vehicle speed. a large radii of curve are required for high speed railway operation. this means that the wear number values are relatively small in compared to conventional operation and it can be assumed that the rail wear in the curve will be less intense than in conventional railway operation. but with increasing vehicle speed and cant deficiency, the values of wear number increase very quickly, as does the specific friction work in wheel/rail contact. this paper primary shows trends of the dynamic vertical wheel force q and the wear number wn values under the specific conditions of high speed railway operation and parameters of multi-body simulations, which are used for evaluation of vehicle/track interaction. the absolute values of these characteristics depend on the design of the railway vehicle and, of course, on the parameters and conditions of the track. acknowledgements this work was supported by the scientific research project of the university of pardubice no. sgs_2020_009. 51 jiří šlapák, tomáš michálek acta polytechnica ctu proceedings references [1] british railway board. permissible track forces for railway vehicles. gm/tt0088, issue 1, revision a, 1993. [2] commision regulation (eu) no 1299/2014 of 18 november 2014 on the technical specification for interoperability relating to the ’infrastructure’ subsystem of the rail system in the european union, 2014. [3] t. zhang, d. huanyun. on the nonlinear dynamics of a high-speed railway vehicle with nonsmooth elements. applied mathematical modelling 76:526-544, 2019. [4] base price wear in the train-path pricing system 2017 instruction for determining vehicle prices. bern: feredal office of transport, 2016. [5] en 14363:2016+a1:2018. railway applications – testing and simulation for the acceptance of running characteristics of railway vehicles – running behaviour and stationary tests. brusel: european committee for standardization, 2018. 52 acta polytechnica ctu proceedings 31:45–52, 2021 1 introduction 2 vehicle/track interaction at higher speeds 3 multi-body simulations 3.1 multi-body model of railway vehicle 3.2 simulation scenarios 3.2.1 influence of running on a straight track 3.2.2 influence of cant deficiency 3.2.3 influence of curve radius 3.2.4 influence of specific conditions around the speed limit of 300 km/h 4 simulation results 4.1 results from straight track 4.2 results from curved tracks 4.2.1 influence of cant deficiency 4.2.2 influence of curve radius 4.2.3 influence of specific conditions around the speed limit of 300 km/h 5 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0181 acta polytechnica ctu proceedings 33:181–187, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague fundamental properties and durability of slope protection spray mortar reinforced with bamboo fibers kazuo fujiyoshia, ∗, takao uedab, masayuki tsukagoshic a kankyo bosai co., ltd., 1-57 akui-cho, tokushima 770-0046, japan b inst. of science and engineering, tokushima university, 2-1 minamijyosanjima-cho, tokushima 770-8506, japan c faculty of engineering, fukuoka university, 8-19-1 nanakuma, jonan-ku, fukuoka 814-0180, japan ∗ corresponding author: k-fujiyoshi@kan-bou.jp abstract. in order to enhance resistance against cracking and durability, short fibers such as steel and organic fibers have been conventionally mixed into spray mortar used for slope protection. in this research, fundamental properties of bamboo-fiber-mixed spray mortar were examined by laboratory tests. the mechanical properties of spray mortar containing bamboo fibers were examined under cyclic wet and dry conditions along with its resistance against freezing and thawing by a spray test. it was confirmed that 0.75% mixture of bamboo fibers in spray mortar most successfully improved mechanical properties and durability. these include adhesion strength to the base surface following exposure to cyclic wet/dry conditions and overall resistance against freezing and thawing. in addition, higher compressive strength and adhesion strength to the base surface were achieved by further mixing in vinylon fibers or fly ash in combination with bamboo fibers. keywords: adhesion strength, bamboo fiber, cyclic wet and dry condition, fly ash, resistance against freezing and thawing, spray mortar, vinylon fiber. 1. introduction the spray application method of mortar for slope protection has been widely used because it can help preserve bedrock conditions from weathering and erosion. however, it has been observed at many sites that sprayed mortar cracks, peels and flakes due to aging. sprayed mortar is directly influenced by the external environment immediately after construction and initial cracking occurs in certain weather conditions. therefore, in recent years, the use of spray mortar mixed with short fibers such as organic and inorganic fibers has been increasing for the purpose of reducing initial cracking and improving flexural toughness [1]. bamboo forests are abundant in japan and have been a familiar source of natural materials. however, proper forest management is somewhat being neglected due partly to the lack of successors who take care of the forests recently. when proper management is neglected, the bamboo forests may trigger landslides and other disasters caused by crowded bamboo growth density and shallow rooting system. as part of the efforts to widening the scope of bamboo material applications, experiments have taken place in the use of short bamboo fibers as a reinforcement medium in concrete and mortar. past experiments [2] have found that bamboo fibers mixed in concrete improve flexural toughness and resistance to crack development. based on these findings, the fundamental properties of bamboo-fiber-mixed-mortar for slope protection were tested in laboratory. then, the durability of bamboo fiber reinforced mortar was tested by spraying on formworks in the actual working spray facilities. the durability examination included two mechanical property tests (for compressive and adhesion strength) under cyclic wet/dry conditions (hereinafter referred to as "cwdc"), and a resistance test in freezing/thawing conditions. 2. experimental program 2.1. properties of fibers the bamboo fibers used in this study were prepared using fibers collected from a bamboo tree with a unique rotating knife. bamboo trees used were anywhere from three to five years old and the cultivar is called moso bamboo grown in awaji, japan. prepared bamboo fibers were then classified by fiber length to l = 20 ± 10 mm using 1.2 − 5.0 mm sieves. the classified bamboo fibers were air dried indoors for about three weeks prior to actual experiments. additionally, commercially available vinylon fibers were mixed into spray mortar to facilitate comparison with spray mortar reinforced with bamboo fibers. the material properties of the bamboo and vinylon fibers are shown in table 1. the dry surface state was defined as a state where the fibers absorbed water for 24 hours and then dehydrated until the fiber surface was dry. to achieve this dry surface state, the fibers that absorbed water were dehydrated for three minutes using 181 https://doi.org/10.14311/app.2022.33.0181 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en k. fujiyoshi, t. ueda, m. tsukagoshi acta polytechnica ctu proceedings a centrifugal dehydrator. bamboo vinylon density: air-dry (g/cm3) 0.77 1.09 density: dried surface (g/cm3) 1.27 1.17 water absorption rate (%) 87.2 12.1 fiber length (mm) 19.8 (av.) 18.0 fiber diameter (mm) 0.59 (av.) 0.2 aspect ratio (−) 34 90 tensile strength (mpa) 154 490 table 1. material properties of bamboo fiber and vinylon fiber. 2.2. laboratory test for fundamental properties 2.2.1. mix proportions of mortar and specimen preparation table 2 shows the mix proportions of mortar for fundamental property tests in laboratory, normal mortar with c = 400 kg/m3, w/c = 55% and s/c = 4.2 was named "n" as a standard composition of spray mortar. based on n, compositions with bamboo fibers added with mixing ratios of 0.25%, 0.75% and 1.25% in volume were each named "b0.25," "b0.75" and "b1.25." in addition, 20% in fine aggregate of b0.75 was replaced with fly ash, which was named "b+fa." in b+fa, w/b was 32% when the binder included fly ash. furthermore, a composition n with 0.75% vinylon fibers was named "v," and the n mortar with 0.375% each of bamboo fibers and vinylon fibers was named "b+v." readily available portland cement (density: 3.16 g/cm3) was used. crushed sand manufactured in naruto, japan was used as fine aggregate (density: 2.57 g/cm3, fineness modulus: 2.63). moreover, type ii fly ash (density: 2.33 g/cm3, specific surface area: 3240 cm2/g, ignition loss: 2.8%) specified under the japanese industrial standard (jis a 6201) was used. the mortar was mixed in accordance with the cement physical test method (jis r 5201), and the fibers were immersed in the mixing water. all compositions in this experiment were stiff consistency mortar with low fluidity close to zero slump. in mortar spraying operations on the field, mortar is sprayed on the bedrock by compressed air generated by a compressor. thus, the sprayed mortar not only adheres to the bedrock but is also compacted. since fundamental property tests are conducted indoors, the specimens were prepared by an elaborate compaction to obtain effects similar to spray compaction without spraying. cylindrical specimens (! 50 × 100 mm) for the compression test were compacted every three layers by tamping down with a thrust bar. 2.2.2. curing and various tests all specimens went through indoor air curing for 28 days (room temperature: 20 ◦c , 75% r.h.), being sealed except for the upper surfaces of the specimens to simulate site conditions. two cases of compression tests were conducted. in the first case, the compressive strength was measured after air curing for 28 days. in the second case, the specimens were subjected to cwdc for 12 days after air curing for 28 days. cwdc were based on the method for the accelerated rock slaking test (jgs 2125) to simulate the environment of natural weathering after spraying on slope surfaces. the specimens were dried at 40 ◦c for 48 hours, then immersed in water at 20 ◦c for 24 hours, and then dried at 110 ◦c for 24 hours. this cycle was repeated three times. a porosity test was performed using cylindrical specimens (! 5 × 10 cm) with a material age of 28 days. the mortar porosity was determined by equation 1, which was consisted of the mass of the specimen dehydrated in a drying furnace at 80 ◦c for 24 hours (w d) and the mass of its surface-dried state immersed in water at 20 ◦c for 48 hours (w s). ε = w s − w d ρ × v × 100 (%) (1) where ε is porosity (%), ρ is density of water and v is volume of specimen. 2.3. spray test using actual plant 2.3.1. mix proportions of mortar and spray method for the spray test, the amount of mixed fibers was determined to be 0.75% as found to perform the best by the laboratory test. five kinds of mortar compositions, n, b, v, b+v and b+fa were selected. it is presumed that the viscosity gets too high due to the influence of the fly ash that replaced fine aggregate in b+fa, and there is a risk of clogging the material hose and nozzle. consequently, the replacement rate of fly ash was adjusted to be 7%, and w/b was 45% when fly ash was used as the binder. mortar spraying was conducted in the facilities normally used for wet spraying. the fibers were mixed in in a small portion at a time to prevent unevenness, and the mixing duration was set to be two minutes per batch. 2.3.2. specimen preparation and various tests figure 1 shows an outline of the formworks sprayed with mortar and the specimens cut out. two types of wooden formworks in the figure were prepared for each mortar composition. three cylindrical specimens (! 50 × 100 mm) for compression test and four specimens for adhesion test were made in a formwork with dimensions of 850 × 300 × 150 mm. in the attempt to take actual spraying on bedrock or existing concrete surfaces into account, two kinds of specimens for the adhesion test were prepared by spraying on rock pieces (! 75 × 25 mm, compressive strength: 207 n/mm2) and on a flat concrete 182 vol. 33/2022 spay mortar reinforced with bamboo fibers fiber type fiber addition rate w/b unit amount [−] [vol. %] [%] ! kg/m3 " c w s fiber pa n − − 55 400 220 1679 0 0 b0.25 bamboo 0.25 55 400 220 1676 1.93 0 b0.75 bamboo 0.75 55 400 220 1670 5.78 0 b1.25 bamboo 1.25 55 400 220 1663 9.63 0 v vinylon 0.75 55 400 220 1662 8.14 0 b+v bamboo, vinylon 0.375, 0.375 55 400 220 1674 2.89, 4.07 0 b+fa bamboo 0.75 32 400 220 1356 5.78 280 table 2. mix proportions of mortar in laboratory test. 850 30 0 300 400 150 specimen for compression test (ij50×100) 25 0 550 specimen for freeze/thaw test (ij100×200) specimen for adhesion test to rock piece (ij75×25) specimen for adhesion test to flat concrete plate unit: mm figure 1. top view of formworks for spraying and sampling of specimens. plate (300 × 300 × 60 mm for interlocking block, flexural strength: 5.5 n/mm2). rock pieces were cut out from sandstone with a stone cutter. the cross section of the stones for bonding with mortar became smoother than the surface of flat concrete plate. moreover, two cylindrical specimens for freeze/thaw test (! 100 × 200 mm) were cut out of mortar sprayed on a wooden formwork with the size of 550×250×250 mm. the sprayed formworks were cured outdoors exposed to natural weather for 28 days from 11/24 to 12/22. during this period, the average temperature was 8.3 ◦c, and it rained for seven days. each specimen except for freeze/thaw test was cut out at a material age of 25 days. when the age reached 28 days, the compression strength test (jis a 1108) and the first case of the adhesion strength test (jscek 561-2013) were conducted. an adhesion strength testing device of building research institute method was used for the adhesion test. after outdoor exposure for 28 days, two specimens for the adhesion strength test were exposed to cwdc for 12 days. then the second case of the adhesion test was performed. three cylindrical specimens (! 50×100 mm) were cut out of mortar after the second adhesion test, and used for the second compression test. 56-day-old specimens were cut out and used for the freeze/thaw test. in this test, 300 cycles of freezing and thawing in water were conducted according to the freeze/thaw test method of concrete (jis a 1148, procedure a), and the relative dynamic modulus of elasticity and the mass reduction rate were obtained. the relative dynamic modulus of elasticity was determined by the ultrasonic propagation velocity with reference to the previous study [3]. 3. fundamental property laboratory test results 3.1. porosity figure 2 shows the results of porosity test at a material age of 28 days. according to this figure, when the mixing amount of bamboo fibers increases from 0% (n) to 1.25% (b1.25), the porosity also increases. this increase is caused by fine bubbles introduced in the mortar with the fibers. the porosity of v mixed with vinylon fibers is the largest. one of the reasons is because the number of vinylon fibers per mortar volume is greater than the number of bamboo fibers. the diameter of vinylon fiber is about 1/3 that of bamboo fiber as shown in table 1. when the same volume of fibers (0.75%) is mixed, the number of vinylon fibers is about 10 times that of bamboo fibers. the porosity of b+v mixed with composite fibers (bamboo/vinylon), is almost the same as b0.75 mixture of bamboo fibers solely. moreover, the porosity of b+fa with the combination of bamboo fibers and fly ash is the smallest because mortar voids are reduced by fly ash. since fine aggregate was replaced with fly ash, voids were filled by the filler effect of fly ash with extra fine particles. in the long term, as the pozzolanic reaction of the fly ash proceeds, the porosity of b+fa will decrease further. 183 k. fujiyoshi, t. ueda, m. tsukagoshi acta polytechnica ctu proceedings 10 12 14 16 po ro si ty (% ) figure 2. porosity of mortar measured in laboratory test. 3.2. compressive strength the results of compressive strength test of 28-day-old specimens are shown in figure 3 "pre wet/dry exposure." the compression strength increased when the addition rate of bamboo fibers increased from 0% (n) to 0.75% (b0.75), but it decreased when the rate further increased from 0.75% (b0.75) to 1.25% (b1.25). in general, the higher the porosity and organic content, the lower the compressive strength. however, when bamboo fiber content increased from 0% to 0.75%, not only the porosity, but also the compressive strength increased. the cause of the increase in compressive strength is inferred to be influenced by the internal curing of bamboo fibers. since the specimens were cured in air, the moisture in the mortar (needed for the hydration reaction) tended to evaporate easily. because moisture contained in bamboo fibers with high water absorption rate (as shown in table 1) is replenished inside the mortar as it dries, the strength of b0.25 and b0.75 is considered to increase from that of n. when bamboo fiber mixing ratio increases to 1.25%, the increased voids (as shown in figure 2) have a more negative effect on the mortar strength. it also becomes difficult to sufficiently disperse the fibers while mixing, so that the compressive strength of b1.25 seemingly decreased. v in which vinylon fibers solely was mixed had a lower compressive strength than n due to its large porosity, and b+v in which bamboo and vinylon fibers were combined had comparable strength to n. b+fa in which bamboo fibers and fly ash were used in combination had the smallest porosity and the highest compressive strength. b0.75 also shows a similarly high-level of strength to b+fa. in terms of strength, approximately 0.75% of bamboo fiber mix ratio proved to be most desirable. when the specimens were exposed to cwdc for 12 days after 28 days of outdoor air curing, the strength of the bamboo-fiber-mixed compositions except b+fa shows a relatively large decrease (see figure 3 "post wet/dry exposure"). the decrease is assumed to be caused by the hollow cells of bamboo 20 30 40 50 c om pr es si ve s tr en gt h (n /m m 2 ) pre wet/dry exposure post wet/dry exposure figure 3. compressive strength of mortar measured in laboratory test. fibers [4]. it is presumed that shrinkage stress is generated in the mortar around the hollow cells while drying in wet/dry cycles, and defects have occurred inside the mortar. the strength of b+fa mixed with fly ash largely increases. since fly ash used, overall amount of the binder increased. this resulted in increased amount of unreacted binder. it is possible that the hydration reaction of the unreacted binder progressed rapidly during water immersion in wet/dry cycles. 4. the results of spray test using actual plant 4.1. porosity figure 4 shows the results of porosity test of 256-dayold specimens (! 50 × 100) cut out in the spray test, together with the porosity of 28-day-old specimens in the laboratory test. according to figure 4, the values of b and v in the spray test are less than those in the laboratory test, and are similar to the value of n. the reason is because the spraying pressure purged the fine bubbles which were entrained while mixing in the fibers [5]. the value of b+fa in the spray test is drastically lower than that in the laboratory test. the decrease was caused by densification of mortar, in addition to bubbles vanished while spraying. since the specimen material age reached as old as 256 days, the pozzolanic reaction of the fly ash progressed. 4.2. compressive strength figure 5 shows the measurement results of compressive strength and density of cylindrical specimens (! 50 × 100 mm) before and after exposure to cyclic wet/dry conditions (hereinafter referred to as "ecwdc"). before ecwdc, the compressive strength of b was similar to that of n and v, while the density of b mixed with bamboo fibers was the smallest. the moisture within the mortar easily evaporates during outdoor curing. in this situation, it is assumed that the bamboo fibers that have absorbed water gradually supplied moisture (the internal curing effect [6]) and the hydration reaction of b further 184 vol. 33/2022 spay mortar reinforced with bamboo fibers 10 12 14 16 n b v b+v b+fa po ro si ty (% ) laboratory test spray test figure 4. porosity of mortar measured in spray test. developed. that is why the compressive strength of b+v is greater than that of v. the reason why b+fa had the largest strength was that the dense mortar matrix was formed by the lower w/b, the filler effect of fly ash and the internal curing of the bamboo fibers. after ecwd, the density of all compositions increased, but the compressive strength of n and b decreased. the decrease was especially drastic in b than in n. the air-dry density of the bamboo fibers is small as shown in table ??, and the inside of the bamboo fiber cells becomes hollow when drying [4]. accordingly, it is considered that the internal defects of the mortar were caused by the stress of drying shrinkage. the compressive strength of v, b+v, and b+fa increased slightly. these increases are attributed to the hydration reaction of the unreacted binder that was promoted during the wet period of cwdc. v and b+v were also mixed with vinylon fibers which do not have a hollow structure like bamboo fibers. as for b+fa, the addition of fly ash resulted in increased amount of powder element that acted as a binding agent, which led to increased compressive strength. the increased compressive strength is assumed to be further improved by the pozzolanic reaction of fly ash over a long period. 4.3. adhesion strength figure 6 shows the measurement results of adhesion strength between spray mortar and rock/concrete surfaces before and after ecwdc. before ecwdc (material age of 28 days), since the rock surfaces cut with a rock cutter were smooth, the adhesion strength with the rock in all compositions was as small as approximately 1.0 n/mm2. peeling fracture also occurred at the adhered interfaces. as for concrete surfaces, the base concrete or spray mortar was destroyed, showing values exceeding 2.0 n/mm2. these values can be regarded as close to the tensile strength of mortar and concrete. regardless of the 2.1 2.12 2.14 2.16 2.18 2.2 20 30 40 50 n b v b+v b+fa d ensity (g/cm 3) c om pr es si ve s tr en gt h (n /m m 2 ) compressive strength: pre wet/dry exposure compressive strength: post wet/dry exposure density: pre wet/dry exposure density: post wet/dry exposure figure 5. compressive strength of mortar measured in spray test. base type, the value of b+v using composite fibers was the largest. adhesion strength after ecwdc decreased on rock/concrete surfaces for all compositions. the decrease was caused by the changes in the mortar volume following ecwdc, increasing the shear strain on the adhered interfaces. for the rock surfaces, peeling fracture occurred at the interfaces. the values of b and b+v were about twice the value of n, and greater than v. for the concrete surfaces, the base concrete or spray mortar was destroyed. the values of b+v and b+fa were larger than other compositions. these results show that compositions with bamboo fibers mixed in have excellent adhesion strength. bamboo fibers inside the mortar absorb water during the wet period of cwdc, and gradually supply moisture in the dry period. it means that drying shrinkage strain on the adhered interfaces are suppressed. 4.4. resistance against freezing/thawing figure 7 shows the mass reduction rate and the relative dynamic modulus of elasticity in the freeze/thaw test results. with respect to the three compositions n, v and b+v, relative dynamic elastic modulus dropped sharply before reaching 200 cycles. these reductions were attributed to the peeling fracture of the mortar pieces at the corner portions of the cylindrical specimens. it is possible that the cracks progressed from the partial sand streaks generated during spraying and led to the peeling fracture. the sand streaks also existed on the specimens of b and b+fa mixed with bamboo fibers, but the relative dynamic elastic modulus showed a gentle decrease, with 89% and 57% at 300 cycles, respectively. after the freeze/thaw test, deterioration due to scaling was observed on the specimen surfaces in all 185 k. fujiyoshi, t. ueda, m. tsukagoshi acta polytechnica ctu proceedings 0 1 2 3 n b v b+v b+fa a dh es io n st re ng th (n /m m 2 ) rock: pre wet/dry exposure rock: post wet/dry exposure concrete: pre wet/dry exposure concrete: post wet/dry exposure figure 6. adhesion strength between sprayed mortar and rock/concrete surfaces measured in spray test. compositions. as shown in figure 7, the mass reduction rate moved momentarily to the negative side in all compositions, but then shifted to the positive side. the cause of the negative shift was that fine cracks occurred at the early stage of the freeze/thaw action, then the mass increased due to the penetration of moisture into the cracks. as the number of cycles increased, scaling or peeling fracture occurred and finally the mass decreased. in the previous study on sprayed concrete [5], it has been reported that helpful air bubbles for freeze/thaw resistance are lost by spraying pressure. when alternative measures were taken to obtain the benefits of bubbles (such as adding hollow microspheres to spray concrete mixed with vinylon fibers), fine cracks were suppressed and reduction in relative dynamic elastic modulus was hardly observed. in this freeze/thaw test, no special measures were taken to compensate for the bubbles lost by spraying pressure. thus, v and b+v mixed with vinylon fibers are considered to have shown small resistance against freezing and thawing. in contrast to n, v and b+v, the freeze/thaw resistance of b and b+fa mixed with bamboo fibers was greatly improved. it can be presumed that bubbles remaining in hollow parenchyma cells of bamboo fibers absorbed and relaxed the expansion pressure during freezing. 5. conclusion in this study, fundamental properties and durability of slope protection spray mortar reinforced with bamboo fibers were experimentally examined. the results and findings are summarized as follows. -1.5 1.5 � ��� ��� ��� m as s re du ct io n ra te (㸣 ) 0 20 40 60 80 100 0 100 200 300 r el at iv e dy na m ic e la st ic m od ul us ( 㸣 ) number of cycles n b v b+v b+fa figure 7. variation curves of mass reduction rate and relative dynamic elastic modulus measured during freeze/thaw test in spray test. 1. as far as strength is concerned, 0.75% mixing ratio of bamboo fibers proved to perform the best. compressive strength decreased following exposure to cyclic wet/dry conditions when only bamboo fivers were mixed. however, when bamboo fibers were mixed with vinylon fibers or fly ash, the decreased strength was improved. 2. concerning spray mortar with 0.75% bamboo fibers mixed in, adhesion strength between the base rock/concrete and spray mortar after exposure to cyclic wet/dry conditions was increased by 180% for the base rock and by 120% for the base concrete, compared to mortar without fibers. the performance improvement was more remarkable when vinylon fibers or fly ash was mixed in in addition to bamboo fibers. 3. spray mortar with 0.75% bamboo fibers had great resistance against deterioration caused by freezing and thawing, and relative dynamic elastic modulus remained at 89% after 300 cycles in the freeze/thaw test (mortar without fibers: 0% at 120 cycles). the mortar combining 0.75% bamboo fibers with fly ash remained at 59% at 300 cycles. acknowledgements we would like to express our gratitude to sugimotoshoten co., ltd. for providing bamboo fiber materials for the research and green system co., ltd. for cooperating in spray test. references [1] s. yan, h. jiao, x. yang, et al. bending properties of short-cut basalt fiber shotcrete in deep soft rock 186 vol. 33/2022 spay mortar reinforced with bamboo fibers roadway. advances in civil engineering 2020:1-9, 2020. https://doi.org/10.1155/2020/5749685. [2] d. k. gupta, r. c. singh. an experimental evaluation of compressive strength and flexural strength of bamboo fiber reinforced concrete. international research journal of engineering and technology 5(6):699-708, 2018. https://www.irjet. net/archives/v5/i9/irjet-v5i9126.pdf. [3] r. takada, s. guo, h. ogata, et al. proc. jp. concr. inst. (kochi) vol 26 pp 1911-6, 2004. [4] j. zhu, h. wang, c. wang. study on the swelling characteristics of bamboo based on its graded hierarchical structure. wood and fiber science 51(3):332-342, 2019. [5] f. taguchi, m. takahashi, n. kishi n, et al. improving the frost resistance of short-fiber-mixed shotcrete by hollow microspheres. 6th international conference on concrete under severe conditions (consec’10), merida, yucatan, mexico, pp 273-8, 2010. [6] v. k. ganesh, k. ramanurthy. the influence of superabsorbent polymer beads used as internal curing agent on the compressive strength of mortar. international research journal of engineering and technology 3(2):410-416, 2016. https://www.irjet. net/archives/v3/i2/irjet-v3i272.pdf. 187 https://doi.org/10.1155/2020/5749685 https://www.irjet.net/archives/v5/i9/irjet-v5i9126.pdf https://www.irjet.net/archives/v3/i2/irjet-v3i272.pdf https://doi.org/10.14311/app.2022.33.0153 acta polytechnica ctu proceedings 33:153–159, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague critical assessment of co2 emission of different concretes: foamed, lightweight aggregate, recycled and ordinary concrete devid fallianoa, ∗, salvatore quattrocchib, dario de domenicob, giuseppe ricciardib, ernesto gugliandoloc a department of structural, geotechnical and building engineering, politecnico di torino, corso duca degli abruzzi, 24, turin 10129, italy b department of engineering, university of messina, contrada di dio, messina 98166, italy c g. gugliandolo s.r.l., via galileo galilei, messina 98100, italy ∗ corresponding author: devid.falliano@polito.it abstract. construction materials contribute to about 75% of the co2 emission of all the construction processes. concrete is one of the most widely used construction materials and is thus primarily responsible for co2 emission. in particular, 8 − 9% of global greenhouse gas (ghg) emission are produced by concrete. co2 emissions can be considerably reduced in the construction phase through a careful selection of materials with low environmental impact or through specific admixtures. in this study, different concretes are taken into consideration, including foamed concrete, lightweight aggregate concrete, recycled concrete and ordinary concrete. a series of mix designs of these four classes of concrete, characterized by a comparable mechanical strength or a comparable density, are taken from the relevant literature and compared to one another in terms of co2 emission. some guidelines or possible research lines aimed at reducing co2 emission are finally outlined in this contribution. keywords: co2 emission, foamed concrete, lightweight concrete. 1. introduction in the last few decades, the attention to environmental issues is more and more increased. the european union issued several guidelines and recommendations in order to reduce global co2 emissions. the objective is to achieve a 40% reduction of the 1990 co2 emission values by 2030 [1]. environmental protection, saving of raw materials and effects of human activities on the climate changes are popular matters that are of interest in diverse social contexts. consequently, the construction industry is also committed to solve these challenges, as witnessed by the large range of research topics in the relevant literature. indeed, the building industry alone is responsible for about 40% of the total energy consumption [2]. more specifically, when coming to the concrete industry, the energy consumption is estimated to be around 3% of the total consumption value, which is associated with an overall co2 emission of about 8-9% of the total co2 emissions [3]. the environmental impact of concrete is due to the fundamental ingredient acting as binder, namely portland cement. indeed, the portland clinker is obtained by cooking a mix of natural and/or artificial soils, including clay, lime, pyrite ash etc. in the blast furnace at 1400◦c. during this preparation phase, the limestone included in such raw materials undergoes the calcination process with significant co2 emission. a twofold environmental impact can thus be recognized in this process: 1) the co2 emission due to the calcination process; 2) the co2 emission due to the energy consumption of the overall process [4]. as a result, the portland clinker is responsible for more than 90% of the overall co2 emission of the concrete industry [5]. based on the above remarks, it appears clear that the most straightforward strategy to mitigate the greenhouse gas (ghg) emission is to reduce the amount of clinker, by partly replacing it with supplementary cementitious materials, such as fly ash and ground granulated blast furnace. the latter technique allows the achievement of different objectives at a time: not only the reduction of the environmental impact in terms of ghg emission by the concrete industry, but also the re-use of by-products and recycled materials, which are converted from negative elements to a useful resource [6]. mineral additions in parallel to clinker portland are allowed by international standards. indeed, in the uni en 197-1 five different classes of cement are reported, depending on the clinker and mineral addition contents. evidently, cements having higher amounts of mineral additions as replacement of portland clinker are characterized by lower co2 emissions. therefore, the co2 emissions due to the concrete industry depend on the type of binder being adopted, which changes from country to country, as the specific culture and the available raw materials influence the preparation phases of concrete. this 153 https://doi.org/10.14311/app.2022.33.0153 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en d. falliano, s. quattrocchi, d. d. domenico et al. acta polytechnica ctu proceedings concept is clearly acknowledged in [7]. besides the introduction of mineral components as partial replacement of the portland cement, which is the conventional strategy adopted in the range of ordinary density concretes, the co2 consumption can be well reduced for special concretes, including foamed concretes (fcs) and aerated autoclaved concretes (aacs). these special concretes are more convenient for the preparation of low-to-medium density concrete mixes. considering the low density, these special mixes are also featured by good thermal insulation properties [8], which is associated with lower energy demands for heating and cooling processes, thus contributing to a lesser amount of co2 emissions not only in the preparation stage, but during the whole service life of the structure [2]. this is particularly convenient in hot climates during the summer season in terms of thermal inertia, when compared to classical insulation panels currently available in the market. recent research on foamed concrete has focused on the determination of the mechanical strength [9] and thermal properties [10], the impact of different ingredients such as foaming agent [11], [12], cement type and presence of superplasticizers [13], the influence of the curing conditions [13], the influence of the presence of by-products and slags, such as fly ash [14] and biochar [15], the influence of mixing intensity [16], the possibility of improving the mechanical characteristics via short fibers [17, 18], or bi-directional grid reinforcement [19, 20] and the fracture behaviour [21–23]. however, to the authors best knowledge, there is no extensive study in the literature concerning the co2 emissions of foamed concrete mixes compared to alternative concrete mixes such as ordinary concretes (ocs), lightweight aggregate concretes (lacs) and recycled concretes (rcs) either by recycled aggregates or by replacing part of the clinker with recycled components. to fill this gap, the present work aims to present a comparative study focused on the co2 emission of such different cementitious pastes. the comparison is made assuming mixes with similar values of mechanical strength or similar values of density. in this way, the evaluation of the co2 emission is made assuming homogeneous conditions among four classes of concrete (ocs, lacs, fcs, racs). a critical discussion is finally made regarding the possibility of adopting alternative concrete mixes (as substitute of ordinary concretes) characterized by lower co2 emissions. 2. materials and mix design in this section, a series of mix designs of different experimental campaigns taken from the relevant literature are presented. these mix designs concern several concrete mixes selected to assess the co2 emissions taking into account different parameters such as density, compressive strength, composition of the mix and main ingredients, presence of mineral additions or recycled elements. through this comparative study, it is possible to shed light on key aspects primarily responsible for the ghg emissions, so as to plan future experimental campaigns on special concretes wherein the reduction of the co2 emission is the main objective. moreover, the sources of the co2 emissions that are inherent in the conventional preparation processes of ordinary and special concretes will be discussed. in particular, the study comprises four different types of concretes, namely fcs, lacs, rcs, and ocs. a preliminary selection process is made in order to identify certain common properties of the different mix designs related to the different types of concrete. as an example, the co2 emission will be comparatively analysed for mix designs characterised by the same compressive strength (but different densities), or alternatively by the same density (but different compressive strengths). the motivation for including these different types of concrete is due to the significantly different compositions in terms of mix design. for instance, in fcs a part of volume is occupied by air voids generated by the presence of a preformed foam, or introduced in the matrix through proper chemical reactions during the initial phases of the preparation. besides these fc specimens, another widely used class of lightweight concretes is represented by lacs. in this case, the cementitious matrix is the same as in ordinary concretes, but the self-weight reduction is achieved by the partial replacement of normal aggregates with lightweight aggregates. in the cases discussed in this paper, expanded clay was used in the specimen preparation. finally, another interesting class of concretes is represented by rcs in which part of the aggregate is replaced by recycled materials. in this specific case, the recycled components are represented by lightweight recycled aggregates as partial replacement of normal aggregates. in this way, lightweight characteristics are achieved also in this class. with regard to fc, three reference studies are included. more specifically, the mix designs presented by some of the authors in [13] and labelled as #1, #2.1 and #4.1 in the referenced paper [13] are included, which are characterized by the presence of cement, water and preformed foam resulting in different densities spanning from about 400 kg/m3 to 800 kg/m3. moreover, another paper presenting mix designs of fcs addressed in this study is [24], where, unlike the previous experimental campaign, the specimens contained also fly ash. also in the latter study the densities span a wide range, from 750 kg/m3 to 1250 kg/m3, so as to supplement the previous experimental campaign [13] with higher-density specimens. finally, the third study is [25], where ultralightweight foamed concrete specimens (density ranging from 100 to 250 kg/m3) with the use of fly ash were documented. in this way, a very broad range of densities of fcs are included in this examination. with regard to lacs, only one benchmark study is considered [26], in particular the specimens labelled 154 vol. 33/2022 co2 emission of different concretes figure 1. distribution of co2 emissions related to 25% of world cement production (data from [7]). lt30, lt40 are included. in this case, the lightweight properties of the specimens were achieved by the addition of expanded clay. in particular, keeping a constant amount of 500 kg/m3 of cement, the authors investigated the effect of different contents of expanded clay ranging from about 200 to 330 kg/m3, obtaining densities from 1280 to 1600 kg/m3 approximately. with regard to rcs, only two specimens were taken from the extensity study in [27], labelled cm20rm and chd50rm. besides the classical aggregates and portland cement, these specimens contained lightweight recycled aggregates and expanded clay. finally, in order to compare the results pertinent to the above-described mixes with traditional concretes having densities of about 2400 kg/m3, two classical mix designs were also included in the study, which are characterized by compressive strengths of around 25 and 30 mpa (comparable to some of the aforementioned special concretes). these specimens have an aggregate-to-cement ratio in the range from 2.7 to 4.78, and a water-to-cement ratio of around 0.5. extensive details of the mix designs of all the considered concretes can be found in the quoted experimental studies, to which the reader is referred to. 3. co2 evaluation a basic procedure for the computation of the co2 is adopted in this study. based on the mix design proportions and the amount of the single components, an estimate of the contribution of each component to the overall co2 emission is performed through assumptions from the literature. in particular, the present study is limited to the assessment of the co2 emission related to the production phase. in other words, this evaluation does not include the emissions related to the transportation from the manufacturing place to the concrete plant, nor the emissions related to the casting in the construction site, nor the lifecycle data (e.g. recovering part of co2 emissions through carbonation of concrete [28]). as a first remark, it is worth noting that the main source of co2 emission in the mix design is represented by the cement. however, slightly different values of the kg co2 emissions per ton of clinker are reported in the literature from country to country [7], as illustrated in figure 1. the differences are due to the different technologies adopted in the various countries in the concrete industry, and due to the different fuels employed. moreover, besides cement, the other predominant ingredients being present in the mix designs analysed in this study are: i) sand; ii) expanded clay; iii) fly ash; iv) water; v) recycled aggregate; vi) foam. these secondary ingredients contribute to the overall co2 emission to a far less extent in comparison with cement. however, they are included for completeness. in particular, the specific co2 emissions of ash and cement are calculated according to the guidelines provided in [29]. whereas the specific co2 emissions of sand, recycled aggregates, water and components of foam are computed according to the instructions provided by the keiti (korea environmental industry & technology institute), environmental declaration office, available at website epd.or.kr/eng/lci/lcico200.do. finally, the specific co2 emissions of the expanded clay are evaluated according to the environmental product declaration provided by the italian company laterlite s.p.a. the calculation of the co2 emissions due to the foam in fcs is made taking into account the different concentration of the foaming agent and the actual density of the foams used in the mix designs documented in the referenced papers. based on the assumptions described above, figure 2 illustrates the co2 emission (expressed in kg of co2 per m3 of concrete) for fcs, lacs and ocs at a comparable compressive strength of 25 mpa and 30 mpa, 155 d. falliano, s. quattrocchi, d. d. domenico et al. acta polytechnica ctu proceedings figure 2. co2 emission comparison alongside related density values among three classes of concrete for equal compressive strength of 25 mpa (left) and 30 mpa (right). figure 3. co2 emission comparison alongside related compressive strength values among three classes of concrete for equal density of 1400 kg/m3 (left) and 1600 kg/m3 (right). figure 4. increase of co2 emission for foamed concrete with and without fly ash in the mix design as a function of the density. 156 vol. 33/2022 co2 emission of different concretes respectively. in the figure, the density of the various classes of concrete is also reported along a secondary vertical axis for completeness, wherein it is easily seen that oc is characterized by a much higher density than both fc and lac. by inspection of the plot, it is seen that for both compressive strength classes (i.e. 25 and 30 mpa), ocs are characterized by a lower quantity of co2 emission in comparison to fcs and lacs. this is justified by the fact that in the mix design of ocs a lower amount of cement is used. indeed, in fcs and in lacs the amount of cement is around 500 kg/m3 in order to attain reasonably good mechanical strengths with lower densities, whereas in the traditional concretes (ocs) typically this quantity does not exceed 350 kg/m3. moreover, lacs are characterized by a slightly higher co2 emission compared to fcs (for equal amount of cement) because the co2 emissions underlying the production process to realize the expanded clay are surely higher than the co2 emissions related to the production of the foam (one order of magnitude higher approximately, based on the previous guidelines adopted for the calculation). considering that the ocs produce lower amount of co2 emissions than lightweight concretes, attention is then paid to the comparison among three classes of lightweight concretes, namely fcs, lacs as well as rcs. figure 3 depicts the comparison of co2 emissions of the three lightweight concretes at equal density of 1400 kg/m3 and 1600 kg/m3, respectively. also in this case a secondary vertical axis is added on the right-hand side of each plot, showing the values of the compressive strengths. based on this comparison, it emerges that rcs are the most favourable ones in terms of both highest compressive strengths and lowest co2 emissions. varying the density, it is confirmed that lacs are characterized by the poorest performance, from both environmental and mechanical perspectives. comparing the two plots on the left and on the right, it is also noticed that the co2 emissions do not vary significantly in relationship to the two analysed values of density. in the case of lacs and rcs, this is due to the fact that the increase of density was obtained by reducing the lightweight aggregates while keeping the cement amount quite constant. analogously, in fcs the increase of density was achieved through a lower amount of foam while keeping the cement amount quite constant. the above statements concerning fcs are valid for medium-to-high densities. nevertheless, by significantly widening the range of the examined densities for fcs it is noted that the density plays a key role in the overall co2 emissions. this is shown in figure 4, which reports the co2 emissions for a range of densities from 150 (ultra-lightweight foamed concretes) up to 1600 kg/m3. in the low-density range, the amount of cement drastically decreases, while the quantity of the foam is simultaneously increased, thus generating a significant reduction of the resulting co2 emissions. values of co2 emissions even five times lower for ultra-lightweight foamed concretes compared to medium-density foamed concretes are indeed observed. however, for lower densities the compressive strength values decrease accordingly, thus making them unable to cope with structural applications [13, 25]. indeed, such ultra-lightweight elements are mostly used in non-structural applications for thermal insulation and acoustic absorption purposes. in figure 4 two sets of data are reported, including mix designs of fcs with and without fly ash. obviously, the fly ash has a considerable effect on the co2 emissions because it is used as partial replacement of the cement in the mix design, thus significantly reducing the co2 emission. as an example, for fcs having density of 750 kg/m3 the addition of fly ash allows the reduction of the co2 emissions of more than 300%. in figure 4 two regression curves for fcs with and without fly ash, respectively, are superimposed. these curves, whose parameters are determined through least-square regression, well agree with the reported or computed values of co2 emission, as confirmed by the high value of the coefficient of determination r2 being higher than 0.98. these curves can be used within the range examined in this paper to predict the co2 emission produced by fcs depending on whether fly ash is incorporated or not in the mix design. 4. discussion and conclusions this study has presented a numerical investigation on the co2 emission produced by four classes of concretes with different mix designs but comparable properties in terms of mechanical strength or density. since the main cause of co2 emission is due to the amount of portland cement, in order to keep the co2 emission to within an acceptably low level, the optimization of the mix design of special concretes plays a key role. in foamed concretes it is well known that abundant use of portland cement is made to obtain good mechanical strengths. therefore, the elaboration of novel mix designs wherein the clinker is replaced by alternative elements, such as fly ash and silica fume, is very useful. however, such replacement is rather limited in practice, because the quantity of cement (despite the introduction of fly ash or silica fume) remains still too high (≥ 500 kg/m3) in most of the studies from the literature. subsequent studies on foamed concretes should be directed toward a compromise between a further reduction of portland cement (e.g. ≤ 400 kg/m3) and good mechanical strengths that can be acceptable for structural purposes (e.g. rck ≥ 20 mpa). an attractive alternative solution is represented by geo-polymer foamed concretes, but relevant literature in this regard is rather scarce. instead, the co2 emissions are notably reduced for ultra-lightweight foamed concretes, in which part of the cement is replaced by fly ash or other mineral additions. similar remarks can be 157 d. falliano, s. quattrocchi, d. d. domenico et al. acta polytechnica ctu proceedings made with regard to lightweight aggregate concretes realized with expanded clay, since also in this case the quantity of portland clinker remains relatively high. in contrast, recycled concretes represent an optimal balance between mechanical strength and co2 emission. evidently, in the latter case the primary aspect to investigate regards the durability of this special kind of concretes. acknowledgements the authors are very grateful to the whitbybird foundation for funding j. a.’s initial work on the original paper [7]; and to the open-oxford-cambridge arts and humanities research council (ahrc grant number ah/r012709/1) doctoral training partnership for funding the continued project. funding for a. m. came from the school of engineering and innovation at the open university references [1] european union 2014 communication from the commission to the european parliament, the council, the european economic and social committee and the 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[29] r. leese, d. casey. embodied co2e of uk cement, additions and cementitious material. report of mineral products association (mpa), fact sheet 18, 2012. 159 https://doi.org/10.1088/1757-899x/615/1/012018 https://doi.org/10.1016/j.proeng.2015.06.157 https://doi.org/10.1016/j.tafmec.2019.102297 https://doi.org/10.1016/j.prostr.2019.08.196 https://doi.org/10.1016/j.conbuildmat.2015.01.051 https://doi.org/10.1016/j.conbuildmat.2012.02.053 https://doi.org/10.1016/j.jclepro.2014.11.015 https://doi.org/10.1016/j.eiar.2014.01.004 https://doi.org/10.14311/app.2022.33.0480 acta polytechnica ctu proceedings 33:480–488, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague microstructural characterization of high-performance steel fiber reinforced geopolymer concrete zoi g. rallia, ∗, stavroula j. pantazopouloua, vladimiros g. papangelakisb a york university, department of civil engineering/lassonde school of engineering, 4700 keele street, toronto, m3j 1p3, canada b university of toronto, department of chemical engineering and applied chemistry, 200 college street toronto, m5s 3e5, canada ∗ corresponding author: zoiralli@yorku.ca abstract. due to growing environmental and economic concerns associated with conventional building materials, research interest gravitates towards the development of novel environmentally friendly materials as alternatives to conventional portland cement concrete. geopolymer concrete is a class of novel advanced and sustainable structural materials that hold promise for the future of infrastructure. its synthesis comprises industrial by-products (fly ash and slag among others) in the role of binder and thus reduces the demand in portland cement leading to a significant carbon footprint reduction. in the present study a high-performance fiber reinforced geopolymer concrete (hpfrgc) is synthesized from first principles and is subsequently characterized, with particular emphasis on its microstructural and mineralogical properties. the study explores the linkage between the microstructure and mineralogy of the precursors, and the properties of the final product. both fresh and hardened hpfrgc are studied. experimental results illustrate the correlation between microstructure, mineralogy and final mechanical properties can be used as an indicator of suitability of industrial by-products for geopolymer precursors. the effect of these choices on stability and physical properties of the material is also explored in the study. keywords: advanced materials, geopolymer concrete, microstructure, mineralogy, sustainability. 1. introduction there is no doubt that nowadays the need for construction building materials has been intensified due to the increase of urbanization rates and the acceleration of global population growth. a commensurate increase in demand for ordinary portland cement (opc) conventional concrete is expected since construction is a basic leverage of economic development and cement is the cornerstone of the modern infrastructure. however, despite of its widespread use, conventional concrete is not a sustainable solution as it impacts the environment with a potential of considerably high cost for mitigating its implications. production of cement, as the main component of opc concrete, is a high-energy consuming process and a major contributor to global warming. note that for every 1 ton of produced opc, an equal amount (i.e., 1 ton) of co2 is generated. additional consumption of energy relating to aggregate production and crushing, contributes to the unsustainable depletion of natural resources. it is expected that the green taxes imposed on cement industry as penalty for not reducing the co2 footprint will skyrocket the cement price. consequently, environmental, and economic concerns emerging from the energy intensive industry of building materials impose a need for reconsideration of every step in their production; one important mitigating strategy is the utilization of innovative alternative cements that have much smaller, or zero co2 footprint. such materials are geopolymer binders, encompassing a variety of industrial wastes in the role of a binder and filler. geopolymers represent the latest advancement in alumino-silicate chemistry and they were first developed by davidovits in the 1970s [1]. alternative binders are used for their production with activation mechanism, chemistry, and composition fundamentally different from that of cementitious materials. the term geopolymers refers to mineral compounds, or blends thereof, consisting of repeated units, such as silico-oxide (-si-o-si-o-) and silico-aluminate (si-o-al-o-) units, generated through a bonding process called geopolymerization. in this process, many small atomic structures, also known as oligomers, are linked together to form a covalently bonded threedimensional network. these geo-chemical syntheses occur in the presence of a reactive aluminosilicate material, also known as precursor, either in alkaline medium in the presence of alkali ions such as na+, k+, li+, ca++ and cs+, or in acidic medium in the 480 https://doi.org/10.14311/app.2022.33.0480 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 microstructure of fiber reinforced geopolymer concrete length diameter aspect ratio density tens. strength elastic modulus lf df lf /df ρf ft ef 13 mm 0.2 mm 81.25 7.9 (g/cm3) 2500 mpa 200 gpa table 1. properties of steel fibres. figure 1. particle size analysis of dry materials. presence of phosphoric and humic acids. due to the high cost of the latter method, the most common route for geopolymer concrete is the alkaline route especially in the presence of sodium or potassium silicate solutions, which are also known as hardeners [2]. generally, geopolymer binders demonstrate superior performance and have great advantages over conventional concrete. the development of strength and is not based on a hydration mechanism, but rather on a rigid three-dimensional network formation and is rapid compared to conventional concrete. contrary to conventional concrete, geopolymer concrete does not possess hydrates in its microstructure; this leads to a higher stability at elevated temperatures [3]. regarding sustainability and life cycle cost, its production is associated with considerable reduction of the carbon footprint as a result of the absence of cement. in addition, the utilization of industrial waste by-products such as fly ash, blast furnace slag or red mud as source materials provides an additional benefit. it adds value to wastes, while the use of such materials instead of disposal is beneficial for the environment. the most critical aspect of developing a geopolymer concrete is the right selection of the aluminosilicate source materials. the selection criteria such as high performance and sustainability are determined through deeper understanding of the physicochemical properties of the material. a suitable source material (i.e., precursor) should be highly amorphous, possess the ability to release aluminium easily and contain sufficient reactive glassy content with low water demand. several studies have been conducted on different types of binders and how they affect the properties of the geopolymer concrete. in literature, a variety of aluminosilicate raw or processed natural materials, such as metakaolin or industrial byproducts such as fly ash, as well as granulated blast furnace slag, have been used alone or in combinations as binders for geopolymer concrete and their mechanical and physical properties along with their microstructure have been studied. despite the growing interest of the research community on the characterization of geopolymer concrete, there is a limited number of investigations that explore the effect of precursor’s properties on the properties of the geopolymer concrete. in the present study, a combination of the most commonly used supplementary cementitious materials, namely metakaolin, fly ash type with cao less than 10% (type f), silica fume and blast furnace slag were used for the production of a geopolymer matrix with enhanced early strength and high performance. industrially produced potassium silicate solution was used as the alkaline reagent, whereas brass coated steel fibres were used for the main reinforcement. subsequently, the geopolymer was cured at ambient temperature and was then characterized in terms of fresh, hardened, physical, microstructural, and mineralogical properties. a relation between the latter and the raw materials used was also identified. 481 z. g. ralli, s, j. pantazopoulou, v. g. papangelakis acta polytechnica ctu proceedings % chem. ffa sf mk-750 ggbfs % chem.compos. ffa sf mk-750 ggbfs sio2 56.5 94.5 52.51 35.2 k2o 0.84 0.59 0.19 0.59 al2o3 23.5 0.41 43.3 8.37 tio2 0.63 < 0.01 1.42 0.34 fe2o3 3.55 0.16 0.47 0.68 p2o5 0.09 0.08 0.06 0.02 mgo 1.08 0.3 0.05 10.7 mno 0.06 0.04 < 0.01 0.27 cao 9.1 0.49 0.08 38.9 cr2o3 0.02 0.02 0.03 0.02 na2o 2.13 0.15 0.35 0.31 v2o5 loi 0.01 1.89 < 0.01 3.5 0.03 2.01 < 0.01 2.01 table 2. chemical composition of precursors. 2. experimental 2.1. materials fly ash type f (ffa), ground granulated blast furnace slag (ggbfs), silica fume (sf), metakaolin750 (mk-750) were used as geopolymer precursors. basalt (bs) and granite (gs) were chosen as partially reactive sands due to their enhanced bonding with the binder [4]. a commercially available potassium silicate with molar ratio (mr) of 1.7 was selected as the hardener (ps) on account of its high performance and stability. preliminary studies conducted prior to this project, had shown that an industrially produced hardener with a fixed mr performs better than a laboratory made hardener. the better performance is in terms of fluidity and repeatability, while temperature changes do not affect its properties. straight steel fibers (f) with length 13 mm and diameter 0.2 mm were utilized for main reinforcement at a 2.5% volume fraction (table 1). finally, a polycarboxylate-based high-range water reducing admixture (hrwra) was used to obtain the desired fluidity. for the best quality control and understanding of the geopolymer concrete, the physical, chemical, and mineralogical properties of the studied precursors were analysed through a series of characterization techniques. 2.1.1. particle size analysis the particle size analysis was performed on a laser diffraction particle size analyser (beckman coulter ls 13 320). for very fine particulate materials, surfactants were used for deagglomeration. figure 1 shows the particle size distribution of the analysed materials. 2.1.2. chemical analysis x-ray fluorescence (xrf) was performed to get the chemical composition of the precursors. the data given in table 2 confirm that the aluminosilicate source materials could be potential candidates as geopolymer binders as they contain elevated amounts of alumina and silica in combination with the low magnesia [5]. however, the cao content in the ffa (close to 10%) poses a concern regarding the speed of setting. 2.1.3. mineralogical analysis the mineralogy of the aluminosilicate source materials is also significant in geopolymerization. for this reason, measurements were made to determine through x-ray diffraction (xrd) the mineralogy and degree of crystallinity of the raw materials. table 3 illustrates the results of the analysis. the presence of mullite and quartz in ffa in combination with the relatively lower amorphous content (compared to the rest) suggests that the combustion temperature of ffa was higher than 850 ◦c but lower than 1500 ◦c [6]. ggbfs was found highly amorphous with the presence of gehlenite to be favourable in terms of geopolymerization. however, previous studies have shown that it cannot be used alone as a geopolymer precursor. during alkalination, gehlenite depolymerizes into hydrates and precipitates alkalis. the resulting product is not a geopolymer, but an alkali activated slag, which despite its high strength, there are problems of stability and durability in case the material comes in contact with water. this is because geopolymerization is incomplete and the alkali cation from the hardener is not chemically bonded to the structure [2]. 2.1.4. reactivity index another reactivity and suitability index was determined through ph measurements. this measurement determines whether there is a risk of flash setting [5]. the ph of aqueous solutions consisting of 5 g of raw material in 50 ml of distilled water was measured after 5 minutes of preparation. generally, values lower than 8 do not indicate danger for instant hardening, whereas fast hardening occurs for values between 8 and 11 and flash-set is highly possible for values above 11 [5]. the recorded values of ph of the materials are listed in table 4. 2.2. methods using the data from the characterization steps and taking into consideration cost and sustainability factors, three different mixtures of geopolymer concrete were studied with different ratios of ggbfs to total binder (0%, 5% and 10% of the precursors). the mixture design (table 5) is based on the fundamental concepts for geopolymerization [7]. 482 vol. 33/2022 microstructure of fiber reinforced geopolymer concrete mineral mineral formula ffa sf mk-750 ggbfs quartz sio2 11.1 1.4 mullite 2al2o3 · sio2 19.4 calcite caco3 0.5 silicon carbide sic 2.8 gehlenite 2cao · al2o3 · sio2 3.2 alite 3cao · sio2 1.2 larnite 2cao · sio2 1.2 amorphous 69.0 97.2 98.6 95.4 table 3. mineralogical composition of precursors in wt.%. ph value after 5 ffa (aq.) sf (aq.) ggbfs (aq.) mk(aq.) 11.5 8.1 11.8 7.1 table 4. ph of aluminosilicate materials. materials (kg/m3): ffa sf mk-750 ggbfs k-silicate h2o basalt granite hrwra mix design solution sand sand gs10 907 145 75 94 355 71 574 203 18 gs5 907 145 80 47 355 71 574 203 18 gs0 907 145 92 355 71 574 203 18 steel fibres: all mixes contained 160 kg/m3 table 5. mixture designs. 3. results & discussion 3.1. fresh properties flowability tests were conducted to evaluate the fresh properties of high performance fiber reinforced geopolymer concrete (hpfrgc). the test was performed in accordance with astm c1856 [8]. the flowability of each mixture was measured immediately after mixing and right before casting. during this measurement, the conical mould of the flow table was filled with a single layer of fresh geopolymer concrete. then, the mould was lifted and after one minute, the reported flow value was calculated as the average of maximum and minimum diameter of the spread. the flow values were 157, 172 and 181 mm for the mixtures gs10, gs5 and gs0, respectively. the increase in flowability with the decrease of ggbfs even in such small proportions was evident and linked to the fineness and high cao content of ggbfs. the final setting time was also measured empirically as the time from the start of casting to the time the geopolymer completely hardens. the final setting times were 15, 22 and 30 min for the mixtures gs10, gs5 and gs0, respectively. the short setting time was attributed to the high ph of the precursors and particularly the high content of cao in both ggbfs and ffa. 3.2. compressive strength a force-controlled compression-testing machine was used to test the 50 mm cubes under compression at a loading rate of 0.259 mpa/s in accordance with astm c39 [9]. to study the age effect on compressive results, the specimens were tested at the age of 1, 7 and 28 days. the compressive strength results of each mixture at different age is presented in figure 2. note that the reduction in the ggbfs did not affect the strength at early ages owing to the substitution by mk-750. however, the absence of ggbfs had a more prominent effect on the final strength of gs0. for the mixing of geopolymer concrete, a 5 l hobart mixer was used. the mixing process comprises of dry mixing of ffa with sf for 3 minutes, followed by the addition of the alkaline reagent which was mixed with water and high-speed mixing (200 rpm) for 10 minutes. next, mk-750 is added and mixed for 5 minutes. the last binder material added was ggbfs and was mixed for another 3 minutes. subsequently the sand mixture was added to the system and the mixing continued for 2 min. at this point hrwra was introduced to lower the viscosity and prepare the mix for fibre impregnation. when the mix gained the desired fluidity, fibres were added and mixing continued for 2 more minutes. the concrete was then casted in cubic moulds of 50 mm size (minimum dimension had to be about 4-5 times the fibre length). mild vibration was then performed through the use of tamping rods. finally, the specimens were 483 z. g. ralli, s, j. pantazopoulou, v. g. papangelakis acta polytechnica ctu proceedings figure 2. compressive strength development of hpfrgc. figure 3. qemscan images of hpfrgc thin sections. (a) phase assemblage maps (b) total mineral contents of each sample by mass. figure 4. schematic of thin section qemscan images and the number of fibres per cross section. 484 vol. 33/2022 microstructure of fiber reinforced geopolymer concrete figure 5. ternary diagram of hpfrgc. sealed with a plastic sheet to prevent water evaporation. the specimens were demoulded after 24 h and cured in water at ambient temperature until the day of testing. 3.3. qemscan analysis in order to study the microstructure and mineralogy of the produced geopolymer concrete a qemscan (quantitative evaluation of materials by scanning electron microscopy) analysis was conducted. the system is widely used in the minerals industry as it provides phase and mineral identification at micrometer-scale. polished thin sections were taken from each hpfrgc cube sample. the qemscan results are illustrated in figure 3a. the resolution of the produced images is 15 µm. figure 3b also shows the quantitative composition of the geopolymer mixtures. the mineral maps of all mixtures contain the same major mineral, k-feldspar. the second prevailing mineral is albite. the results agree with what is expected from a geopolymerization process as the k-feldspar and albite are frameworks with 3-dimensional networks. more specifically the produced concrete is a k, na-poly(sialate-disiloxo) with atomic ratio si/al=3 whereas its geological analogue is sanidine and albite [6]. the white parts of the images represent the porosity. the blue parts correspond to fibres with their brass coating intact while the black represent the steel fibres that have lost their coating during hardening. 3.4. fiber distribution qemscan images are used as a tool to interpret the distribution of fibres in the thin sections. a schematic (figure 4) is used to assess the number of fibres crossing an arbitrary plane, a parameter that is critical for assessing their contribution to the mechanical properties of the solid material. the particles with black and blue colour having a dimension in their short axis at least equal to the fibre diameter df = 0.2 mm were identified as fibres (the typical fibre is a cylinder with a diameter of 0.2 mm oriented randomly with respect to the plane of the cross section, therefore in general the intersection has an elliptical shape). in a random arrangement of fibres, the number of fibres, nf , crossing a unit area through hpfrgc is estimated from the cross-sectional area of a single fibre having a diameter (df ) and the volumetric ratio of the fibres, vf [10]. nf = 0.64 vf d 2f (1) equation 1 has been derived assuming only 50% of the fibers to be effective on account of random spatial orientation; for preferential arrangement of the fibers effected by the casting procedure this factor may be higher. for vf = 2.5% and df = 0.2 mm, the nf should be 0.4 per mm2. based on the aforementioned data, ideally the number of fibers at the given randomly cut 11 × 11 mm cross-sections should have been around 54; on average the three samples shown in figure 5 agree quite well with this estimation. 3.5. ternary mix design the microstructure of the studied system was assessed by a ternary diagram (figure 5), plotted from qemscan of the most relevant minerals reported in 485 z. g. ralli, s, j. pantazopoulou, v. g. papangelakis acta polytechnica ctu proceedings figure 6. (a) porosity distribution in hpfrgc measured with qemscan analysis ; (b) fractured hpgc specimen with bright core at 28 days. figure 3b. the corners of the triangle correspond to 100% stoichiometric content of the main oxides, i.e., sio2, al2o3, k2o, na2o & cao. the sides represent axes, depicting weight fractions that range from 0 to 1. any point within the triangle has coordinates on the system of the three axes that represent the fractional ratio of mass content of the standardised oxides; the combination characterizes the stability of physical and durability performance of the resulting composite and therefore may be used to distinguish products that could be classified as alkali activated materials (aam) from geopolymers. in the present study all the systems show very similar phase assemblage which is relatively close to that of the stoichiometric geopolymer [7]. 3.6. porosity qemscan images were further used to provide a rough estimation of porosity and its distribution at the analysed thin sections. figure 6 shows the porosity as a function of size and surface area. the porosity of the hpfrgc can be attributed to three different mechanisms: (a) the production of gas, (b) the fast increase of viscosity with time during casting, and (c) the compaction of the material resulting from simple casting rather than extrusion [11] the first mechanism is a result of incorporation of sf in the mix design. it is worth noting that studies have shown that even small content (< 1 %) of elemental silicon in sf could generate gas approximately 1 to 11 ml/gr of sf [12]. more specifically, this elemental silicon is oxidized in an alkaline environment (equation 2). h2 is generated as a result of water reduction and acts as a foaming agent (equation 3). the overall reaction forms si(oh)4 as shown in equation 4. the chemically induced porosity seems to yield micropores in the range of 0 − 50 µm. this is clearly shown in figure 6 where the porosity distribution is depicted demonstrating that all mixtures have similar percentages of 0 − 50 µm micropores. si0 → si4+ + 4e− (2) 4h2o + 4−e → 2h2 + 4oh − (3) 4h2o + si0 → 2h2 + si(oh)4 (4) the increased viscosity due to the addition of very fine and high-density particles such as sf, mk-750 and ggbfs and the incorporation of the fibres also contributes to the pore formation as the diffusion of the generated gas is enhanced [11]. finally, compaction also affects the porosity, as it depends on the viscosity of the mixture. in case of g10s, the mixture with the highest amount of ggfbs was difficult to compact as much as the other two. this is evident from the high content of pores with maximum size greater then 500 µm in g10s, as shown in figure 6(a). owing to the micropores generated as described above the geopolymer materials are generally lightweight with an apparent density lower than 2 kg/l. for example, the self-weight of the three mixtures tested was 1.9 kg/l. after testing the cubes in compression, and upon examination of fragments from the core of the failed specimens, a whiteish 486 vol. 33/2022 microstructure of fiber reinforced geopolymer concrete discoloration was observed. since this phenomenon could not be studied thoroughly in the hpfrgc broken cubes, as they maintained their integrity on account of the embedded fibres, a series of unreinforced high performance geopolymer concrete cubes were tested in compression to study their fracture surface. although the specimens were cured in water, the core of the specimen is brighter than the rest because it is dried. this is illustrated in figure 6(b). the extent of the dried core area is related to the age of testing, as at the early ages is relatively small while it increases with the time. the reaction of geopolymerization starts from the core and expands through the volume of the material. in the white regions of the core the pores have smaller size compared to the dark regions and with the expansion of the white region (development of geopolymerization) the porosity decreases in the specimen. with the progress of geopolymerization the size of pores becomes gradually smaller, whereas water is expelled as hardening propagates. 4. conclusions from the above experimental investigation, the following conclusions are drawn. 1. a full characterization in terms of physical, chemical, and mineral properties is needed to evaluate the suitability of locally available aluminosilicate source materials as geopolymer binders. the determination of the main precursor should be determined taking into consideration the results of the characterization tests. 2. the phase assemblage of the hpfrgc indicates successful geopolymerization similar to natural rocks such as k-feldspar and albite. the produced hpfrgc is k,na poly(sialate-disiloxo). 3. examining the fresh and hardened properties of hpfrgc, it is concluded that ffa with cao content close to 10% can be used as main precursor in a geopolymer concrete that can harden in ambient temperature and yield high compressive strengths. however, this is accompanied by progressively shorter setting times as ggbfs is added to enhance the final strength. ggbfs has a negative effect in viscosity too, and thus handling of the material. finally, the studied hpfrgc is an excellent candidate for cases where early strengths are needed, as the geopolymer concrete developed in 24 h after casting almost half strength it would have developed in 28 days. 4. all the three mixtures have acceptable number of fibers as proven by examining randomly cut sections. 5. the porosity of the studied hpfrgc is induced by chemical reactions, viscosity, and consolidation. for the chemically induced porosity, pretreatment of aluminosilicate materials to eliminate the free silicon is needed, while workability or viscosity modifying admixtures could be used to decrease the porosity generated from high viscosity and poor consolidation. acknowledgements the authors duly acknowledge the contributions of dr. grammatikopoulos and sgs, canada in carrying out the characterization and microstructural analysis. the support of wöellner, and drain bros. excavating ltd. by donating the alkaline reagent and the basalt and granite sands, respectively, is also acknowledged. references [1] j. davidovits. solid phase synthesis of a mineral blockpolymer by low temperature polycondensation of alumino-silicate polymers, 1976. [2] z. g. ralli, s. j. pantazopoulou. state of the art on geopolymer concrete. international journal of structural integrity 12(4):511-33, 2020. https://doi.org/10.1108/ijsi-05-2020-0050. [3] [4] a. reggiani. types of automatic mixing systems for geopolymer mortar/concrete production and 3d printing geopolymer camp 9-11 july 2019, saint quentin, france, 2019. [5] j. davidovits. geopolymer chemistry & applications, 4th ed, geopolymer institute, saint quentin, france, 2015. [6] j. davidovits, m.izquierdo, x.querol, et al. the european research project geoash: geopolymer cement based on european coal fly ashes, technical paper 22, geopolymer institute library, 2014. https://www.geopolymer.org/wp-content/uploads/g eoash.pdf. [7] w. m. kriven. geopolymer-based composites, ceramic and carbon matrix composites, comprehensive composite materials ii,ăelsevier, oxford, 8(5): 269-279, 2018. https: //doi.org/10.1016/b978-0-12-803581-8.09995-1. [8] american society for testing and materials. c1856/c1856m: fabricating and testing specimens of ultra-high-performance concrete, 2017. [9] american society for testing and materials (astm). c39: standard test method for compressive strength, 2005. [10] n. d. archontas, s. j. pantazopoulou. microstructural behavior and mechanics of nano-modified cementitious materials adv. concr. constr, technopress, 3(1):15-37, 2015. doi:https://doi.org/10.12989/.2015..3.014. [11] e. prud’homme, p. michaud, e. joussein, et al. silica fume as porogent agent in geo-materials at low temperature. journal of the european ceramic society 30(7):1641-8, 2010. https: //doi.org/10.1016/j.jeurceramsoc.2010.01.014. 487 https://doi.org/10.1108/ijsi-05-2020-0050 https://www.geopolymer.org/wp-content/uploads/geoash.pdf https://doi.org/10.1016/b978-0-12-803581-8.09995-1 https://doi.org/10.12989/.2015..3.014 https://doi.org/10.1016/j.jeurceramsoc.2010.01.014 z. g. ralli, s, j. pantazopoulou, v. g. papangelakis acta polytechnica ctu proceedings [12] m.-h.zhang, v. m. malhotra, j. wolsiefer. determination of free silicon content in silica fume and its effect on volume of gas released from mortars incorporating silica fume. aci materials journal 97(5), 2000. https: //www.researchgate.net/profile/min-hong-zhang/p ublication/250613568_determination_of_free_sili con_content_in_silica_fume_and_its_effect_on_v olume_of_gas_released_from_mortars_incorporati ng_silica_fume/links/5dcff98492851c382f43f4e9/ determination-of-free-silicon-content-in-silic a-fume-and-its-effect-on-volume-of-gas-release d-from-mortars-incorporating-silica-fume.pdf. 488 https://www.researchgate.net/profile/min-hong-zhang/publication/250613568_determination_of_free_silicon_content_in_silica_fume_and_its_effect_on_volume_of_gas_released_from_mortars_incorporating_silica_fume/links/5dcff98492851c382f43f4e9/determination-of-free-silicon-content-in-silica-fume-and-its-effect-on-volume-of-gas-released-from-mortars-incorporating-silica-fume.pdf https://doi.org/10.14311/app.2022.33.0518 acta polytechnica ctu proceedings 33:518–524, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague life cycle assessment of carbon concrete composites: a circular economy path beyond climate mitigation? christoph scopea, ∗, edeltraud guentherb, torsten mieleckec, julia schuetzc, eric muendecked, konstantin schultzed a technische universität dresden, chair of sustainability management and environmental accounting, 01187 dresden, germany b unu-flores, unu institute for integrated management of material fluxes and of resources, 01067 dresden, germany c lcee life cycle engineeering experts gmbh, 64295 darmstadt, germany d ingenierbüro grassl gmbh, 13158 berlin, germany ∗ corresponding author: christoph.scope@tu-dresden.de abstract. sustainable construction and materials play an ever-important role to stay within our planetary boundaries. in support, innovative carbon concrete composites (ccc) promise significant raw material savings by integral design. we aim to illustrate current environmental hotspots and a feasible recycling scenario of ccc that meets circularity requirements. we modelled a cradle-tograve life cycle assessment for two potential building structural applications (sandwich wall, ceiling reinforcement) made of ccc. we based our recycling scenario on previously conducted large-scale experiments. results show a relative larger energy intensity and abiotic depletion of fossil fuels for variants of ccc but lower global warming. yet, recycling is, second to embodied emissions of basic materials, the driving force of total environmental impacts. the presented recycling path (demolition, pyrolysis for carbon fabric, reuse in fiber fleece) offers less "green credentials" than steel. keywords: carbon concrete composite, life cycle assessment, recycling. 1. introduction the european construction industry including buildings contributes to 9% in eu’s gdp, but is also held accountable for roughly 90% of used mineral resources, 40% of energy consumption, and 35% carbon emissions [1]. to minimize negative effects on our planetary boundaries, a dedicated pathway to sustainable construction is key. an ever-increasing regulatory attention is paid to climate change and circular economy that could eventually improve sustainable development goals #9 to #13. circular economy, as synthesized by potting et al. [1], allows to align environmental strategies along nine "rs". decisionmakers are asking to what extend new building materials can help achieving these goals. they may offer options of considerable raw material savings ("r2 reduce") by integral design ("r5 refurbish"), remedial maintenance ("r0 refuse") or closing material loops ("r8 recycle"). in this study, carbon concrete composites (cccs) are examined as type of textile reinforced concrete (trc). ccc is understood as cement-based carbon fiber reinforced composites using a fine-grained concrete [2] and mesh-like textile reinforcements or carbon fiber rods as basic materials. in comparison to steel reinforced concrete (src), the use of highperformance carbon fibers offers a much wider "range of application specific properties" [3]. cccs still need to showcase whether its application is environmentally, socially and financially viable [4] though seemingly functionally superior including beneficial raw material savings. despite of over 20 years of intensive studies on ccc by two german collaborative research centers in dresden and aachen [5] and a succeeding joint research center "c3 carbon concrete composite" little is known about sustainability of ccc. an initial scoping review revealed that pertinent literature on ccc is dominated by structural and mechanical tests. there is only a few published work on sustainability assessments. many cases refer to carbon fiber reinforced plastics (cfrps) in other product contexts, e.g. automotive or railway. for example, das [10] applied lca for cfrp components in cars quantifying primary energy consumption (pe) and greenhouse gas emissions to show lightweight advantages. cfrp has also been studied for reinforcements of bridges and façades. cadenazzi et al. [15] assessed five life cycle impact assessment (lcia) indicators whereas hájek et al. [8] calculated four, e.g. global warming (gw) and acidification (ap): cfrp performed better than src in both cases. for ccc, williams portal et al. [9] examined reinforcement technologies, e.g. src and ccc, with cradle-to-gate lca. ccc scored the lowest total environmental impact. laiblová et al. [7] compared 518 https://doi.org/10.14311/app.2022.33.0518 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 lca of carbon concrete composites materials value* unit life cycle examined product source carbon concrete composites (ccc) 34-109 [kg co2eq/m2] a1-c3** façade sandwich panels [6] 21.83 [kg co2eq/m2] a1-c3 façade element [7] 0.35*** [kg co2eq/mpa] a1-a3 façade element [8] carbon fiber reinforced plastics (cfrp) 295.93 [kg co2eq/m3] a1-a3 slab, case "ca1" [9] 345 [kg co2eq/kg] a1-a3 car component, "pan smc" [10] 30-52 [kg co2eq/kg] a1-a3 car component, various manuf. [11] cfrp rebar/strip 19.70 [kg co2eq/kg] a1-a3 rebar [12] carbon fiber 17.43-34.39 [kg co2eq/kg] a1-a3 pan carbon fiber [13] 24.20-31 [kg co2eq/kg] a1-a3 lignin and pan carbon fiber [10] 11.40 [kg co2eq/kg] a1-a3 pan carbon fiber [14] epoxy resin 5.25-8.25 [kg co2eq/kg] a1-a3 gabi database, for de/europe steel reinforced concrete (src) 129 [kg co2eq/m2] a1-c3 façade sandwich panel [6] 26.33 [kg co2eq/m2] a1-c3 façade element [7] concrete c25/30 197 [kg co2eq/m 3] a1-a3 german mix, cement 11.1% epd 0.09 [kg co2eq/kg] a1-a3 ∼ epd reinforcing steel 0.31 [kg co2eq/kg] a1-a3 girders for concrete slabs epd0.68 [kg co2eq/kg] a1-a3 reinforcement steel wire epd * all indicator values are based on dimensions given in each study. ** "life cycle" refers to examined life cycle stages using nomenclature of en 15804:2014. *** normalized to tensile strength in bending table 1. reported values for global warming within past studies (sources given upon request). façade elements made of src, ccc, and trc with basalt or glass fibers by cradle-to-grave lca. they identified no clear winner, though all trc variants did better in 4 out of 6 lcia indicators as compared to src. hülsmeier et al. [6] showed in a cradle-to-grave lca that ccc performed better than src on gw (-13 kg co2eq/m2). a multifunctional, vacuum-insulated façade element made of ccc even outperformed that standard panel of ccc (-74 kg co2eq/m2). recently, stoiber et al. [12] used epd data to analyze bridge designs made of ccc, src, and steel. again, ccc showed lower gw (-34 to -48%), but higher ap (-40 to 48%) scores. those studies served as benchmark, but were partly limited in data availability, e.g. for carbon fiber fabric or end-of-life (eol). to address these research gaps, we compared variants of a prefabricated sandwich wall system and strengthening schemes for an existing concrete ceiling made of either ccc or src. our research aims to identify hotspots for carbon concrete from an environmental perspective and is guided by following: 1. what environmental impacts are characterizing compared structural applications? 2. what level of energy and resource depletion is embodied in each material life cycle? 3. is substitution of src with carbon concrete an eligible measure to foster circular economy? we improved the analysis by collecting nondisclosed company data, updated materials for ccc, and integrated data from joint research projects, e.g. recycling processes. for practitioners, the results signal current hotspots along the value chain of carbon concrete to proactively start redesign strategies. civil engineers will be enabled to discuss material choices within the early design process to improve environmental sustainability. for city planners, we illustrate indicators allowing to measure policy impact and scale. for scholars, we will add insights to existing research dialogues in the area of ccc and environmental sustainability. the method and data used in this article is described in section 2. results and discussion are presented in section 3. section 4 concludes upon our research objectives. 2. materials and methods we began with a scoping systematic review to maintain evidence-based research practices and identify comparable datasets. table 1 illustrates a current 519 ch. scope, e. guenther, t. mielecke et al. acta polytechnica ctu proceedings sandwich wall system ceiling slab reinforcement variant a1 (src, c25/30) variant b1 (ccc, prefab.) variant c (src, c25/30) variant d (ccc) unit inner / outer shell / height 0.19 / 0.08 0.14 / 0.03 0.047 0.006 m xps insulation 0.035 na m n.x (outer shell) 15 na kn load capacity na 5 kn/m2 u-value 0.27 na w/(m·k) sound insulation 77 71 na db concrete 25/30 (% in-situ) 601.7 (40%) na 109.8 (100%) na kg/m2 special concrete for ccc na 406.05 na 14.1 kg/m2 steel reinforcement b500 46.31 na 2.99 or 0 na kg/m2 stainless steel b500 ng 1.04 na 0 or 2.99 na kg/m2 carbon fiber fabric na 2.16 na 0.65 kg/m2 polystyrene 2.7 na kg/m2 net weight 651.75 410.91 112.99 14.75 kg/m2 table 2. functional unit for two structural applications and life cycle inventory (own illustration). lack of coherent knowledge: the range of reported values for just gw varies a lot. we see the importance of well-defined functional units: differences on level of basic materials, e.g. carbon fiber versus steel girders (up to +9, 118%), are substantially greater than on level of entire structures, e.g. ccc versus src used in façade panels (-15 to 73%). 2.1. functional unit and life cycle inventory our reference functional unit was defined as 1 m2 of: a) a prefabricated sandwich wall made of src for a 3-storied building or b) a ceiling slab reinforcement of an existing src slab to increase load capacity from 1.5 to 5 kn/m2. each functional unit is not solely mass-based, but fulfils rated structural and mechanical requirements (see table 2). the wall systems were designed as non-stiffening sandwich outer walls with total dimensions 6 × 2.8 meters. we did not consider connections or cut-outs for existing openings, facades or corners. for the wall systems, we modelled four variants: two variants (a1-a2) of src with different concrete compressive strengths (c25/30 and c50/60) both using in part in-situ concrete. the two variants made of ccc (b1b2) were dimensioned to meet the same requirements. variant b1 was modelled as entirely prefabricated, the other b2 required partly in-situ concrete. the existing ceiling slab was 23.5 cm thick using c25/30. for the ceiling reinforcement, one src and ccc variant each (c and d) was calculated while assessing only the reinforcement materials. calculated material input data in table 2 reveals: the special concrete for ccc amounts to 67.4%, respectively 12.8% of concrete c25/30 in variant a1 and c. weight of steel reinforcement amounts to 4602,192% of carbon fiber fabric. total net weight differs by 240.84 (wall) or 98.05 kg/m2 (ceiling). 2.2. system boundaries and cut-offs our cradle-to-grave lcas encompassed all life cycle stages except the use stage. we followed the nomenclature of en 15804 for life cycle stages. service life was assumed 50 years as suggested by german regulatory bodies. no extended service life scenario was applied. petroleum-based polyacrylonitrile (pan) was used as precursor for ccc. for use stage, no conversion scenarios, replacement cycles or maintenance activities were expected. component layers subject to wear, e.g. paint coats, were not considered, too. the base scenario for eol: the wall is demolished, the building materials are separated and transported to dedicated plants for further processing. based on findings of joint research center "c3 carbon concrete composite", we assumed that 95 percent of initial carbon fiber fabric was reclaimed for material reuse after processing by pyrolysis. material reuse was modelled as producing glass fleece with recycled carbon fibers. avoided production of primary glass fleece was credited. we also modelled "r8 recycle" paths for concrete fraction, (stainless) steel, and insulation materials by road gravel, recycled steel, and thermal recovery. for cut-offs, we neglected plants, machines, and infrastructure necessary for manufacturing and construction. the lcas were mainly set-up for reference area germany. for base scenario, transport distances for the materials ranged between 50 and 500 km. we used gabi software version 9.2.0.58 sp 38 and cml2015 as lcia method. the dataset for carbon fiber rovings and a special fine-grained concrete mix "c3b2-hf-2-145-5" similar to c80/95 was based on another research subproject [16] within joint research 520 vol. 33/2022 lca of carbon concrete composites center "c3 carbon concrete composite". we used non-disclosed data of a manufacturer of carbon fiber fabrics and concrete plants otherwise generic datasets of gabi. ���� �� ��� ��� ��� ��� ���� &uhglw�iru�qdwxudo�jdv��khdw� :dwhu 6xoskxulf�dflg (sr[\�uhvlq 7udqvsruw 1lwurjhq 3urfhvv�vwhdp��qdwxudo�jdv� :dvwh�lq�pdqxidfwxulqj 3rzhu�pl[ 3$1�ilehuv figure 1. hotspot analysis cradle-to-gate for carbon fiber fabric (own illustration). 3. results and discussion 3.1. hotspot analysis for single life cycle stages first, we discuss hotspots identified within manufacturing and upstream activities (cradle-to-gate). figure 1 shows results for the carbon fiber fabric. the manufacturing of precursor material and succeeding production processes (summarized as "pan fibers") are the dominating influencing factors. only odp (ozone depletion) and partly adpel (abiotic depletion, non-fossil resources) are also strongly influenced by electricity consumed and its underlying power mix. manufacturing companies, e.g. toray, hexcel, sgl carbon or teijin, should thus concentrate on corporate efficiency strategies within production and identify suitable green alternatives to natural gas. all other value chain members should increase the share of renewables within their power mix [11] to refer to circular economy strategy "r2 reduce" [1]. moreover, that industry cluster should foster research on alternatives to petroleum-based pan like lignin (see section 3.3) assorted to "r0 refuse" [1]. for concrete, hotspot analyses are already published [17] and not displayed here. ���� ���� ���� ���� �� ��� ��� ��� ��� ���� '��&uhglw�iru�jodvv�iohhfh &���3\uro\vlv &�����'hprolwlrq�dqg�wudqvsruw figure 2. hotspot analysis of the end-of-life stage for carbon concrete (own illustration). second, we look at the eol stage in our base scenario. this includes demolition, transportation to pyrolysis plant, pyrolysis [18], manufacturing of a recycled carbon fiber fleece (a case of downcycling). figure 2 shows a dissatisfying picture: pyrolysis is the dominating process due to its energy intensity. except for indicator ap, none of the seven other environmental impact categories can be offset. the crediting along that "r8 recycle" [1] pathway for the glass fleece production is not enough. thus, current recycling technologies for carbon fiber fabric in ccc offers a true circular economy path, but has a negative net environmental effect, e.g. it requires more energy input than credited. although better than facing difficulties like wind turbines [19], this is far from ideal. 3.2. cradle-to-grave comparison results for the full life cycle illustrate a mixed picture. exemplary, we show adpfos (abiotic depletion, fossil resources), pe, and gw scores for each considered life cycle stage (see figure 3-5). for the wall system, each square meter of variant b1 made of ccc leads to lower carbon emissions (gw: -6.1 kg co2eq/m2) but is characterized by higher energy intensity (pe: +238 mj/m2) and fossil resource depletion (adpfos: +263 mj/m2). same is true for ceiling reinforcement: -2.6 kg co2eq/m2, +94 and +74 mj/m2. apparently, the value chain of carbon fiber and its fabric uses low-carbon power generation as compared to steel. base scenario shows another 521 ch. scope, e. guenther, t. mielecke et al. acta polytechnica ctu proceedings � ���� � ��� ��� ��� ��� ����� ����� 0 -� p ð 9duldqw�$���65&� 9duldqw�%���&&&� figure 3. comparing adpfos along all life cycle stages for wall systems (own illustration). comparative advantage for src with respect to recycling (pe: -108 mj/m2) and resulting credits (-36 mj/m2). about 85% of used steel derives from recycled and remelted material in german construction industry. this is an ongoing success story of a circular "r8 recycling" [1] path already paying back for src. for indicators not shown here: ccc variant b1 keeps an advantage on ap and eutrophication (ep); odp is roughly equal; src variant a1 has lower damage in photochemical ozone creation (pocp). figure 5 illustrates examined effects of four scenarios (violet bars): positive values imply an environmental advantage of variant b1 (ccc); negative values vice versa a disadvantage. for base scenario, variant b1 has lower gw scores (up to +11.9 kg co2eq/m2) than all other design variants (blue bars). scenario 1 assumed a 20% share of steel imported from china within the src value chain. this increased all lcia indicators negatively for variant a1 (gw: +16.9). scenario 2 modelled real location sites of carbon fiber and fabric manufacturers. the advantage of low-carbon water power generation in the us and scotland is entirely offset by environmental impacts of the resulting long-distance oversea transportation (gw: -1.1). scenario 3 ignored the eol stage (gw: +14.8) and scenario 4 assumed a future reuse of recycled carbon fibers for the same purpose, namely carbon fiber fabric (gw: +16.4). this is a clear signal to foster that true recycling path. sensitivity analyses tested for data uncertainty (orange bars): a dataset for steel reinforcement with -50% lower and a dataset for carbon fibers with +50% higher carbon emissions is chosen. the advantage for gw turns negative (gw: -7.8 / -6.8). ��� � �� �� ��� ��� ��� 0 -� p ð 9duldqw�&��65&� 9duldqw�'��&&&� figure 4. comparing pe scores for ceiling slab reinforcements (own illustration). 3.3. limitations as with any lca, (model) choices were made with effects on overall results. first, we did not quantify the use of lignin-based pan, though there is promising research that illustrates lignin as future, more sustainable sourcing. other alternatives to high purity pan, e.g. based on algae or waste cotton linter are rather at an experimental stage. second, we did not study natural fiber materials, e.g. flax, jute, or sisal. they still lack a good balance of impact strength, stiffness, and toughness within the desired flexural and tensile properties. both strategies would support "r0 refuse" [1] and diminish the volume of petroleum-based pan and geopolitical risks. third, we did not explicitly model strategies towards green concrete. we used an optimized special fine-grained concrete mix that was developed and tested for ccc [16]. the benchmarking src was modelled with german average mix of c25/30 and c50/60 by the german cement works association. fourth, we did not account for recycled aggregate concrete or other types with recycled materials. as this is a feasible option for both, src and ccc, so we did not add this complexity to "r8 recycling" [1]. currently, crushed concrete was credited for road gravel. fifth, future analyses could evaluate eco-designs related to modularity and removable joints. this path would follow circular economy strategy "r3 reuse" [1]. sixth, we do not include extended service life scenarios, yet. 4. conclusions our study aimed to earmark comparative environmental advantages for the innovative building material carbon concrete and existing circular economy potentials. we applied cradle-to-grave lca compar522 vol. 33/2022 lca of carbon concrete composites ��� �� � � �� �� �� )xoo�olih�f\foh nj�&2�ht�pð ¨�9duldqw�$��%���zdoo� ¨�9duldqw�$��%� ¨�9duldqw�%��%� ¨�6fhqdulr����$��%�� ¨�6fhqdulr����$��%�� ¨�6fhqdulr����$��%�� ¨�6fhqdulr����$��%�� ¨�6hqvlwlylw\�6whho�uhlqi���$��%�� ¨�6hqvlwlylw\�&duerq�ilehu��$��%�� figure 5. difference in gw scores between wall system variants a (src) and b (ccc) for full life cycle of design variants, scenarios, and sensitivity analyses (own illustration). ing src and ccc variants. scope et al. [20] may meet expectations of the scholarly community framing lca as main pillar of a life cycle sustainability assessment. first, we identified hotspots of environmental impacts located in manufacturing stage of carbon fiber fabric and concrete as well as during recycling processes of ccc. second, we quantified abiotic depletion and primary energy consumption for each material life cycle for a functional unit of 1 m2. during manufacturing and construction stages, the examined wall system variant a1 made of src is less energyintensive (adpfos: -181 mj/m2; pe: -102 mj/m2), but emits roughly +14% more greenhouse gases. for the full life cycle, both materials get closer on gw (+6%), but drift apart on adpfos and pe (-24%/19%) in base scenario. analogies are observed for the ceiling slab reinforcement: src keeps an advantage on adpfos and pe (-53/-54%), but emits +2.7 kg co2eq/m2. net weight of ccc structures is significantly reduced by 59% (wall system) and 665% (ceiling). third, we have discussed a couple of realized or potential circular economy paths. currently, tradeoffs exists: using variant b1 for a whole 3-storied building could amount to increased primary energy of +107,957 mj, but achieved savings of 109 tons materials and 2.8 tons greenhouse gases emissions. city planners then need additional sustainability criteria. overall, our assessment indicates that ccc provides a circular economy material loop, less material weight, and resulting lower carbon emissions. carbon concrete may improve the climate mitigation potential of buildings, but industry and building professionals are encouraged to emphasize discussed technology pathways to improve energy intensity and specific circular economy strategies. we extensively discussed our study’s limitations with room for further research in many fields. practitioners shall not "overlook" [9] carbon concrete but study the new material and start piloting projects to empirically test estimated savings. acknowledgements the authors thank the german federal ministry of education and research for funding the program "twenty20 partnership for innovation" and the entailed project "c3 carbon concrete composite" (grant no. 03zz0320a). references [1] abergel t, dean b, dulac j, hamilton i, wheeler t 2018 2018 global status report: towards a zero-emission, efficient, and resilient buildings and construction sector (iea and unep). 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[15] t. cadenazzi, g. dotelli, m. rossini, et al. life-cycle cost and life-cycle assessment analysis at the design stage of a fiber-reinforced polymer-reinforced concrete bridge in florida. advances in civil engineering materials 8(2), 2019. https://doi.org/10.1520/acem20180113. [16] k. schneider, m. butler, v. mechtcherine. carbon concrete composites c3 nachhaltige bindemittel und betone für die zukunft. betonund stahlbetonbau 112(12):784-94, 2017. https://doi.org/10.1002/best.201700058. [17] j. de brito, r. kurda. the past and future of sustainable concrete: a critical review and new strategies on cement-based materials. journal of cleaner production 281, 2021. https://doi.org/10.1016/j.jclepro.2020.123558. 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[20] c. scope, e. guenther, j. schütz, et al. aiming for life cycle sustainability assessment of cement-based omposites: a trend study for wall systems of carbon concrete: dresden nexus conference 2020 session 4-circular economy for building with secondary construction materials to minimise resource use and land use. civil engineering design 2(5-6):143-58, 2020. https://doi.org/10.1002/cend.202000024. 524 https://doi.org/10.4028/www.scientific.net/amr.1106.69 https://doi.org/10.1061/(asce)mt.1943-5533.0001160 https://doi.org/10.1007/s11367-011-0264-z https://doi.org/10.1007/s11740-018-0802-7 https://doi.org/10.1016/j.jclepro.2020.124300 https://doi.org/10.1520/acem20180113 https://doi.org/10.1002/best.201700058 https://doi.org/10.1016/j.jclepro.2020.123558 https://doi.org/10.1021/acssuschemeng.8b01026 https://doi.org/10.1016/j.scitotenv.2019.135054 https://doi.org/10.1002/cend.202000024 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2021.31.0018 acta polytechnica ctu proceedings 31:18–26, 2021 © 2021 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague possibilities of high-speed railway turnout data description adam hlubuček czech technical university in prague, faculty of transportation sciences, konviktská 20, prague, czech republic correspondence: hlubuada@fd.cvut.cz abstract. this paper aims to summarize possibilities how to create data models of high-speed railway turnouts. the turnouts designed for high-speed operation require specific geometric solution. as the uic railtopomodel is considered an international recommendation in the field of data modeling of railway infrastructure, this issue is assessed in terms of models based on its principles. solving this problem can affect the future development of the multipurpose railway infrastructure model gradually emerging at the ctu railway laboratory in prague using the railtopomodel principles. whereas the railtopomodel itself does not define any specific types of entities, the railml® 3.1 specifications are also used for assessment purposes. turnouts are viewed both in terms of topology and in terms of functional infrastructure. in the final sections, recommendations are given on how to deal with the problems found, e. g. in terms of implementation into the multipurpose railway infrastructure model. keywords: high-speed turnout, railway infrastructure, railtopomodel, railml. 1. introduction a railway turnout is a very complex technical equipment consisting of many components. it allows the vehicle movement to be transferred between two continuing tracks. conventional turnouts can be simply described by a not very large set of parameters, as the radius of each of their branches has a constant value. in order to increase speed in the branching direction, the high-speed turnouts are designed using a more complex geometric solution. this is because they require precise track geometry adapted to limit the high dynamic forces generated by the passage of a vehicle. [1, 2] in order to fulfill multiple needs of many specific process information systems (for example automatic train control systems, automatic train operation systems and train scheduling applications), an exact and unambiguous data description of railway infrastructure is necessary. detailed knowledge of the track geometry parameters is especially essential to calculate the static speed profile, which defines the allowed speed for certain types of railway vehicles all along the route. the ctu railway laboratory in prague focuses to support development of the above-mentioned technologies. [3] among others, the multipurpose railway infrastructure model is being developed at the workplace. the data model is designed to create the infrastructure core database, which structure should enable to meet the needs of related applications and allow to export the data to standardized formats (for example railml® 3). this paper proposes several possible ways how to implement the high-speed turnouts to the data model. 2. selected specifics of high-speed turnouts maximum allowed running speed of a railway vehicle when passing a horizontal curve is limited by the value of unbalanced lateral acceleration (caused by superelevation deficiency) and its rate of change with respect to time. abrupt change of superelevation deficiency and the rate of change of superelevation deficiency as a function of time are often used to express these characteristics. if no superelevation is applied to a segment of a track (or to a branch of a turnout) the speed essentially depends on the radius of curvature and the way it changes. [4] in order to avoid abrupt change of superelevation deficiency, transition curves are usually provided between a straight and curved segment of a track. their purpose is to change the radius from infinity at the straight segment of the track to the radius of the following circular arc gradually, i. e. without abrupt change of curvature. transition curves can be used to join two circular arcs with different radius of curvature as well. nowadays, clothoids are mostly adopted to provide transition curves. the clothoid is a curve with a linear variation of curvature. in the case of a superelevated arc, gradient due to superelevation (also called as superelevation ramp) is often applied within the adjacent transition curve. [5] the most of ordinary turnouts are designed without using transition curves and without superelevation. it causes relatively high values of abrupt change of su18 https://doi.org/10.14311/app.2021.31.0018 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 31/2021 possibilities of high-speed railway turnout data description figure 1. the course of curvature in the turning branch of the j60-1:33,5-8000/4000/14000-phs turnout perelevation deficiency affecting the passing vehicle. in order to reduce lateral forces between wheel and rail while allowing high-speed train operation, it is necessary to adjust the turnout geometric parameters accordingly. high running speed requires high value of radius which significantly affects the shape and size of a turnout. therefore, slim turnouts adapted for high-speed operation usually reach the designed length of several hundred meters. the radius of the branch arc typically reaches a value of several thousand meters also in the turning direction, in such cases. for this reason, long switch rails are necessary for vehicle to be able to turn onto another track. these switch rails often require several point machines to be handled, then. an additional point machine may be required to manipulate the movable parts of the frog. [6] in some cases, the geometry of a high-speed turnout branch is optimized by the means of transition curves application. for example the turnout type designated as j60-1:33,5-8000/4000/14000-phs (which is the first turnout with unstable curvature developed in the czech republic) is designed using two clothoid curves in the turning branch. turning away from the straight direction, the branch begins at the radius of curvature of 8000 m, gradually changing to 4000 m (at the length of 38 m), which is the radius of the following circular arc (of the length of 40 m), followed by the second transition curve (54 m long, in this case) ensuring change the radius of curvature to 14 000 m. the course of curvature in the turning branch of this turnout is shown in the figure 1. continuation of this clothoid can ensure that the inflection point is reached at an appropriate distance so that it allows the construction of a simple crossover assembled from these turnouts with axis spacing of 4,75 m. [7] 3. railway infrastructure data description based on the uic railtopomodel the railtopomodel initiative was introduced in 2013 [8] in order to create a common generic standard as regards the railway infrastructure data modeling. in 2016, the railtopomodel was released as uic international railway solution 30100 [9]. the uic railtopomodel [10] represents a topological model of railway infrastructure based on a „connexity graph“. the whole model consists of several subsystems, as they follow: • base • topology • positioning • net entity • location [11] the model makes it possible to describe the railway infrastructure at several different levels of detail. socalled net elements and net relations are used to create the topology of a network. individual net elements can be grouped in order to express an element at a less detailed level. for functional reasons, the most detailed level should be defined with the use of linear elements, which can represent, for example, tracks or line sections (depending on the intended purpose of the specific model use case). it is possible to express certain position within a net element using intrinsic coordinate. this coordinate can take value from 0 to 19 adam hlubuček acta polytechnica ctu proceedings figure 2. the base, network and topology modules of the multipurpose railway infrastructure model 1, whereas 0 means the beginning and 1 indicates the end of a linear net element. two net elements can be joined only at the position of 0 or 1, which is realized by the means of net relations. the navigability attribute defines whether is enabled to pass through the given net relation or not (resp. in which direction). the railtopomodel also allows to define several positioning systems of two types. it is possible to create both linear and geometric positioning systems. linear and geometric coordinates belonging to these positioning systems can be assigned to individual intrinsic coordinates of net elements as well as to appropriate locations of net entities. net entities themselves represent different types of railway infrastructure objects and characteristics which can be localized to the topological layer. the railtopomodel itself do not define any specific types of net entities. these types can be introduced as individual extensions to the model. depending on the nature of localized object or property (eventually level of detail), the net entities can be placed using spot, linear and area type of location. a location of a net entity can be created based on both intrinsic and external (linear or geometric) coordinates. several different locations for the same net entity instance are enabled. linear and area locations can cover several (sections of) net elements which is provided by the associated net element class. in some cases, as regards these locations, it is essential to specify values of the keeps orientation and application direction attributes. [9, 11] currently, the most advanced use case of the railtopomodel approach is apparently the railml® 3 data format, [12] as mentioned in more detail further. 20 vol. 31/2021 possibilities of high-speed railway turnout data description figure 3. the positioning system, net entity and location modules of the multipurpose railway infrastructure model 4. multipurpose railway infrastructure model development in order to deploy infrastructure data repository for several key projects dealing with spatial aspects of the railway domain, the ctu railway laboratory in prague decided to implement the core model, largely based on the railtopomodel principles. the multipurpose railway infrastructure model is a relational data model which directly allows it to be used as a database meeting mysql standards. at the same time, however, it reflects an object-oriented approach to railway data modeling. each type of data object is divided into several tables connected by identifying relations representing successive specializations of the particular object types. by simply joining such interconnected tables, a comprehensive overview of individual specialized objects can be obtained. in addition, this approach allows to easily add, remove and change the attributes common for several different types of objects. whenever possible, the relational model tables reflect the railtopomodel classes in terms of naming, attributes and connections. the base, network and topology modules (as can be seen in the figure 2) basically correspond to the railtopomodel layout, from this point of view. for functional purposes, the cardinality of n:n has been introduced, as regards some relations. in such cases, appropriate interconnection tables are added to the model to express the assignment. in the case of more substantial functional changes, new names of the tables are used in order to avoid confusion. 21 adam hlubuček acta polytechnica ctu proceedings figure 4. the visualization module of the multipurpose railway infrastructure model such changes are primarily about implementation of the associated coordinate and associated net element classes converted to the associated point and associated section tables and adjustment of the concept of the net entity locations which distinguishes between physical and functional (associated) location of net entities. for technical reasons, the mechanism of assigning linear a geometric positioning system coordinates to associated points and physical locations has also been modified. these issues are related to the positioning system, net entity and location modules shown in the figure 3. inspired by the railml® 3.1 specifications, [12] the concept of net entity templates has been added to the the multipurpose railway infrastructure model. the net entity table has been adapted for this purpose. the table is located at the interface between the core and use case-specific part of the model. the model has also implemented the visualization module introduced by the railml® 3.1 data format. in addition to linear and geometric positioning systems, it also allows a user to create screen positioning systems and visualization objects. the visualization module implementation can be seen in the figure 4. the use case-specific part includes concrete types of net entities representing various objects and properties which can be localized to the topological layer. the relevant data classes are specializations of the net entity class allocated within the core model at the interface with the use case-specific part. in order to ensure feasible data export and import, it is possible to design these classes and their attributes taking into account the structure of the railml® 3 exchange format. it is appropriate to focus on the geometry, functional infrastructure and protectively physical assets (not yet defined in version 3.1) substructures of the infrastructure subschema. 5. railml® 3 infrastructure subschema railml® [13] is an open source xml-based data format used in the railway branch. the railway markup langue was originally designed to enable data exchange in the scope of timetabling. later on, the specifications [14] were extended to the subsystems of rolling stocks, infrastructure and interlocking. in recent years, attention has been focused especially on the third version development. as regards the infrastructure schema, railml® 3 is essentially based on the railtopomodel principles, whereas the implemented core model is extended to specific aspects mainly related to the net entity subsystem. the first public version of railml® 3 was released on 19 february 2019. [12] noting the fact that railml® 3 respects the use case-oriented approach, it can be said that railml® 3.1 takes into account the following use cases: • network statement • schematic track plan • interlocking module engineering data • simulation [15] each of the above listed use cases is related to either the infrastructure or the interlocking schema or both. depending on further implementation of other use cases, the particular schemas will be complemented with new aspects, especially new specific types of net entities and their attributes. regarding the railml® 3 file structure, the infrastructure schema is represented by the infrastructure element contained in the top-level railml element. the infrastructure element can contain several other subelements representing particular thematic views, as stated below: • topology • geometry • functional infrastructure • physical infrastructure • infrastructure states • infrastructure visualizations [12] the allowable structure has been gradually evolving during the the railml® 3 development and was established as described further. the topology module is 22 vol. 31/2021 possibilities of high-speed railway turnout data description essentially based on the railtopomodel topology subsystem. the corresponding topology element includes container elements for net elements, net relations and networks. their subelements represent specific instances of the respective types of topological objects. all available information about net entities is contained in the geometry and functionalinfrastructure elements, as the physicalinfrastructure element does not yet have internal structure defined. these modules also use the pattern of container elements with nested specific instances of net entities of its type. the geometry module consists of horizontal curves, gradient curves and geometry points. these track geometry objects are specializations of the net entity class as defined in the railtopomodel specification. therefore, they can be localized to the topology layer and using coordinates of positioning systems. another group of net entities introduced by the railml® 3.1 infrastructure schema is represented by the functional infrastructure module containing various railway infrastructure facilities and characteristic (e. g. lines, operation points, tracks, switches, crossings, signals, platforms, speeds and track gauges) from a functional point of view. that makes it probably the most complex part of so far evolved railml® 3. next to the functional infrastructure, the module of physical facilities (in the sense of assets) should be developed at a later date, depending on the upcoming use cases. the physicalinfrastructure element does not yet have internal structure defined. furthermore, railml® 3.1 is able to express infrastructure states in relation to individual elements. that makes it more powerful tool than the current version of railtopomodel which can only define one period of validity, as regards selected objects. the last of the railml® 3.1 infrastructure modules deals with infrastructure visualizations. this module enables the infrastructure to be visualized by making it possible to display various objects on the screen using screen coordinates. [12] 6. possible representation of turnouts using the railtopomodel and railml® 3.1 modeling principles a turnout – as the subject of data description – can be viewed from several different points of view. as regards the railtopomodel and railml® 3 methodics, we can distinguish between the topological and the net entity approach. concerning the railml® 3.1 version, it can be described either as part of a functional infrastructure or as an interlocking device. prospectively, the physical infrastructure perspective should also be implemented. [9, 12] 6.1. topological approach regarding the topological point of view, an ordinary turnout can be modeled using three interconnected linear net elements representing the tracks adjacent to the turnout. there is no net element to represent the turnout itself, at this level of detail. the turnout is only implicitly expressed from the functional perspective with the use of net relations connecting the net elements (see the subfigure a) of the figure 5). each net relations connects two net elements in the roles a and b. each net element can be connected at the point with an intrinsic coordinate of the value of 0 or 1. for each net relation, the attribute of navigability must be defined. the navigability attribute can reach the value of ab, ba, both or none expressing the direction in which the net relation is navigable. [9, 12] for modeling the ordinary turnout as described above, two net relations with navigability of the value of both and one net relation with navigability of the value of none are used. it is also possible to express the turnout as a separate non-linear net element connected to three linear net elements. this kind of description can facilitate the localization process of the net entity expressing the appropriate technical equipment of the turnout. nevertheless, there is no possibility to express the navigability attribute in an appropriate manner, as it cannot be distinguished in which direction to continue (or not) after reaching the non-linear net element (see the subfigure b) of the figure 5). the possible solution is to consider the turnout as a linear net element (see the subfigure c) of the figure 5) or to proceed to an ever more detailed level in order to distinguish the individual branches of the turnout to be considered as two separate net elements (see the subfigure d) of the figure 5). the approach of two separate branches allows to easily model directional conditions and other detailed features related to the turnout (especially in the case it is not desirable to express them for the switch as a whole). nevertheless, the necessary number of net elements and net relations increases. in general, the railtopomodel net elements are supposed to be dimensionless. [9] nevertheless, in terms of some use cases, e. g. as regards the railml® 3 format, the need to introduce the length attribute for the net element instances has arisen. [12] in order to assign the value of the length attribute to each net element correctly and precisely, it is desirable to disintegrate the turnout non-linear net element to three separated linear net elements interconnected at one specific point (see the subfigure e) of the figure 5). the internal topology of every single turnout is similar to the first said topological example, then. in addition, there are several possible interpretations of how his model could be explained. the specific point of interconnection may represent the point of the frog, the intersection of the continuous and branching track axis, the point of the track axis at 23 adam hlubuček acta polytechnica ctu proceedings figure 5. possible expressions of an ordinary turnout using net elements and net relations the level of the tip of tongue etc. each of these specific representations has its significance for a particular use case and affects in which way the attribute values are assigned to particular net elements. in particular, this approach makes it possible to create predefined and easily repeatable patterns of internal topological representation of turnouts. 6.2. net entity approach while the topological point of view allows to express turnouts using net elements and net relations, the perspective of net entities makes it possible to describe them with factual parameters and to locate them to the topological layer and using external coordinates. as already mentioned, the railtopomodel specifications itself state no specific types of net entities, as this is a matter of particular use cases. [9] regarding the railml® 3.1 infrastructure schema, there is a switchis complex type defined in order to describe a turnout from the functional perspective. it serves to express the permissible structure of individual switchis elements, nested in the switchesis container element, being a part of the functional infrastructure view. together with the crossing, bufferstop and border complex types, the switchis complex type is one of the specializations of the tracknode complex type derived from the functionalinfrastructureentity complex type. further, this is a specialization of the entityis complex type which is derived from the complex type based on the locatednetentity railtopomodel class. [12] the “is” substring was added to the designation of the complex type in order to distinguish between the infrastructure and interlocking net entity. for this time, we focus on the switch entity as part of the functional infrastructure view, accordingly. the switchis complex type is designed to express all applicable types of turnouts, i. e. including the slip turnouts. it may have the following attributes defined: • basedontemplate • belongstoparent • branchcourse • continuecourse • type [12] the type attribute allows to distinguish between the ordinary switch, outside curved switch, inside curved switch, single switch crossing and double switch crossing constructions. the basedontemplate and belongstoparent attributes are intended to contain a reference to other elements of the switchis complex type (in the role of the template or parent in relation to the described instance). the branchcourse and continuecourse attributes can take the value of left or right expressing the direction of the particular branches. 24 vol. 31/2021 possibilities of high-speed railway turnout data description these turnout branches can be described as separate elements nested in the switchis element. the leftbranch and rightbranch elements are relevant for an ordinary (resp. outside curved or inside curved) switch. the straightbranch and turningbranch elements can be used up to twice within any switchis instance expressing either of the switch crossing types. all the above-mentioned elements expressing any of the turnout branches are of the switchcrossingbranch complex type. [12] for this reason, they can be described by the following attributes: • branchingspeed • joiningspeed • length • netrelationref • radius [12] the branchingspeed and joiningspeed attributes indicate the permitted speed of the described turnout branch in kilometers per hour with respect to the direction of travel (which may be relevant, for example, in the case of some types of ordinary turnouts). the length attribute expresses the length of the turnout branch in meters. the netrelationref attribute allows the turnout branch to be assigned to one of the net relations from the topological layer. finally, the radius attribute means the radius of the arc of the relevant turnout branch expressed in meters. alternatively, the directional conditions of turnout branches can be modeled independently of the relevant switchis elements, using the horizontalcurve complex type from the geometry view. the net entity representing a horizontal curve can be described by the curvetype, azimuth, deltaazimuth, radius and length attributes. a similar approach can also be used to describe the permitted speed in the switch branch. the speed can be expressed using the speedsection elements with the possibility of nesting an unlimited amount of elements referring to related speed profiles. next to the maximum speed, the speedsection complex type makes it possible to express, if the described speed limit is temporary and signaled. [12] 7. assessment of the usability of the railtopomodel and railml® 3.1 modeling principles for the needs of high-speed turnouts description as regards the topological point of view, there is no difference between high-speed turnout and conventional railway turnout, if it is an ordinary turnout. depending on the desired level of detail and functional requirements, any of the above-mentioned approaches to modeling the railway network topology can be useful. if necessary, the branches of high-speed turnouts can be modeled in even more detailed level, which the approach of net elements connected by net relations seamlessly allows. these net elements of more detailed levels can be standardly aggregated in order to create the less detailed ones. the only possible difference from the current state of the railtopomodel and railml® 3.1 data format may lie in the fact that new values of the descriptionlevel attribute should be introduced if the more detailed topological description needed. only the macro, meso and micro values are currently available, while the macro and meso levels are not sufficiently detailed to show turnouts. although the irs 30100 document [9] also outlines the nano level, it has not been clearly specified and included among the usual levels yet. expressing high-speed turnouts as net entities (if we focus on the functional infrastructure view) is more problematic. the railtopomodel itself is not in charge to express so specific types of objects such as turnouts at all. [9] although the railml ® 3.1 format provides a relatively extensive and sophisticated apparatus for describing the basic characteristics of railway turnout, [12] it is still not sufficient to express all the geometric properties of high-speed turnouts properly. some turnouts designed for high-speed operation, as seen in the the j60-1:33,5-8000/4000/14000phs turnout type example, have a relatively complicated geometric solution of the branches. the single radius attribute is far from sufficient to describe all directional conditions. it would be necessary to define several new attributes to express required properties of the above-mentioned turnout type. its turning branch consist of three different curves, two of which are transition curves with unstable curvature. although such parameterization of directional conditions would be feasible, it would never be able to cover all possible geometric solutions of turnout branches (including those that may be developed in the future). many of these parameters would be completely unnecessary when describing most other turnouts. a more progressive approach is to extend the allowable internal structure of the switchcrossingbranch complex type elements by introducing nested elements expressing subsections of a turnout branch. each such subsection could represent a segment describable by one curve type. such segments could be described by similar attributes as the horizontalcurve elements. these elements make it possible to distinguish between different curve types, such as straight, arc and clothoid. [12] nevertheless, a significant disadvantage lies is the fact that it is not possible to assign the radius of the arc at different positions of the horizontal curve directly. this problem arises especially when describing transitions curves. for this reason, it would be appropriate at least to replace the radius attribute of the horizontalcurve complex type with two similar attributes expressing the radius at the start point and at the end point of the described curve. 25 adam hlubuček acta polytechnica ctu proceedings another solution for one of the railml® 3 upcoming subversions is not to describe turnouts in terms of geometry within the switchis elements at all and use the description of directional conditions exclusively through the horizontalcurve elements instead. this solution would require modifying the horizontal curve description structure similar to the previous proposal. in addition, it would be appropriate to provide the possibility to assign individual horizontal curve to a specific turnout branch. as another point, the turnout branch entity itself could only be described using an aggregated data item of minimal radius, as this is critical in many respects. in cases where this would be sufficient, such a value could be used as the only descriptive aspect of the geometry, if so described. it is always a matter of a specific use case as accurate and detailed description is required. 8. conclusions this article provided insight into the specifics of highspeed railway turnouts and the possibilities of their data modeling. when compared with the principles of the railtopomodel and railml ® 3.1 data format, it was found that there is no significant difference between a conventional ordinary turnout data model and a possible high-speed turnout model, as regards the topological point of view. nevertheless, the functional infrastructure interpretation of a turnout – as offered by the railml ® 3.1 infrastructure model – is insufficient, in this respect. some high-speed turnouts require the geometry to be modeled in a much more precise way. if it is decided to include such turnouts in the multipurpose railway infrastructure model, there are several ways to implement it. we can either increase the number of attributes or extend the internal structure of the turnout data description or solve the geometry of the turnout separately and reference it using external links. these approaches would allow an accurate and detailed description of a high-speed turnout. it is also possible to use only an aggregated data item expressing the value of the minimum radius in each turnout branch instead of the detailed description of its geometry. such a simplified approach is appropriate especially in less demanding use cases. references [1] m. hamadache, s. dutta, o. olaby, et al. on the fault detection and diagnosis of railway switch and crossing systems: an overview. applied sciences 9, 2019. https://doi.org/10.3390/app9235129. [2] m. hamarat, s. kaewunruen, m. papaelias. a life-cycle cost analysis of railway turnouts exposed to climate uncertainties. iop conference series: materials science and engineering 471, 2019. https://doi.org/10.1088/1757-899x/471/6/062026. [3] m. leso, d. kamenický, p. koutecký, a. hlubuček. dopravní sál fd – železniční laboratoř pro výuku i výzkum. vts správy železnic 1, 2019. [4] m. kapitsa, d. bobyr, a. desiak. determining permissible speed of tilting train in curved track. science and transport progress bulletin of dnipropetrovsk national university of railway transport 70: 29–40, 2017. https://doi.org/10.15802/stp2017/109537. [5] m. lindahl. track geometry for high-speed railways. tech. rep., royal institute of technology, stockholm, 2001. [6] h. ossberger. modern turnout technology for high speed [online]. https://www.unece.org/fileadmin/ dam/trans/doc/2016/ter/heinz_ossberger_modern_ turnout_technology.pdf. [7] l. raif, b. puda, j. havlík, m. smolka. design of high-speed turnouts and crossings. iop conference series: materials science and engineering 236, 2017. https://doi.org/10.1088/1757-899x/236/1/012044. [8] uic. feasibility study uic railtopomodel and data exchange format [pdf], 2013. http://www.railtopomodel.org/files/download/ railtopomodel/270913_trafit_ finalreportfeasibilitystudyrailtopomodel.pdf. [9] uic. railtopomodel v1.0 [pdf], 2016. http://www.railtopomodel.org/en/download/ irs30100-apr16-7594bca1524e14224d0.html?file= files/download/railtopomodel/180416_uic_ irs30100.pdf. [10] railtopomodel. railtopomodel [online], ©2020. http://www.railtopomodel.org. [11] uic. railtopomodel v1.1 [online], 2017. https://uic.org/rail-system/railtopomodel. [12] railml. railml schema version 3.1 [xsd], 2019. [13] railml. railml [online], ©2002–2020. https://www.railml.org. [14] railml. the railml subschemas [online], ©2002–2020. https://www.railml.org/en/user/subschemes.html. [15] railml. use cases [online], ©2002–2020. https://www.railml.org/en/user/use-cases.html. 26 https://doi.org/10.3390/app9235129 https://doi.org/10.1088/1757-899x/471/6/062026 https://doi.org/10.15802/stp2017/109537 https://www.unece.org/fileadmin/dam/trans/doc/2016/ter/heinz_ossberger_modern_turnout_technology.pdf https://www.unece.org/fileadmin/dam/trans/doc/2016/ter/heinz_ossberger_modern_turnout_technology.pdf https://www.unece.org/fileadmin/dam/trans/doc/2016/ter/heinz_ossberger_modern_turnout_technology.pdf https://doi.org/10.1088/1757-899x/236/1/012044 http://www.railtopomodel.org/files/download/railtopomodel/270913_trafit_finalreportfeasibilitystudyrailtopomodel.pdf http://www.railtopomodel.org/files/download/railtopomodel/270913_trafit_finalreportfeasibilitystudyrailtopomodel.pdf http://www.railtopomodel.org/files/download/railtopomodel/270913_trafit_finalreportfeasibilitystudyrailtopomodel.pdf http://www.railtopomodel.org/en/download/irs30100-apr16-7594bca1524e14224d0.html?file=files/download/railtopomodel/180416_uic_irs30100.pdf http://www.railtopomodel.org/en/download/irs30100-apr16-7594bca1524e14224d0.html?file=files/download/railtopomodel/180416_uic_irs30100.pdf http://www.railtopomodel.org/en/download/irs30100-apr16-7594bca1524e14224d0.html?file=files/download/railtopomodel/180416_uic_irs30100.pdf http://www.railtopomodel.org/en/download/irs30100-apr16-7594bca1524e14224d0.html?file=files/download/railtopomodel/180416_uic_irs30100.pdf http://www.railtopomodel.org https://uic.org/rail-system/railtopomodel https://www.railml.org https://www.railml.org/en/user/subschemes.html https://www.railml.org/en/user/use-cases.html acta polytechnica ctu proceedings 31:18–26, 2021 1 introduction 2 selected specifics of high-speed turnouts 3 railway infrastructure data description based on the uic railtopomodel 4 multipurpose railway infrastructure model development 5 railml® 3 infrastructure subschema 6 possible representation of turnouts using the railtopomodel and railml® 3.1 modeling principles 6.1 topological approach 6.2 net entity approach 7 assessment of the usability of the railtopomodel and railml® 3.1 modeling principles for the needs of high-speed turnouts description 8 conclusions references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0100 acta polytechnica ctu proceedings 26:100–106, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app design of tuned mass dampers for large structures using modal analysis jan štěpánek∗, jiří máca czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jan.stepanek@fsv.cvut.cz abstract. tuned mass damper is a device, which can be highly useful when dealing with excessive vibration and is widely used in many engineering fields. however, its proper design and optimization is a complicated task. this study uses mode superposition method to speed up the evaluation of dynamic response. the speed of response calculation allows for a quick calculation of frequency response function and numerical optimization of tuned mass dampers. this optimization method is demonstrated on a numerical example of a cable stayed footbridge. the example compares a simplified and widely used design method of tuned mass damper with numerical optimization. keywords: dynamic response, modal coordinates, tuned mass damper, optimization. 1. introduction the tuned mass damper (tmd) is a device composed of a spring, mass and a viscous damper, which is widely used in order to reduce mechanical vibration. however, its high efficiency is conditioned by appropriate design. unfortunately, the closed form solution for its optimal parameters has been found only for special, simplified cases. numerous studies dealing with optimal design of tmd have been published since the beginning of the 20th cetury, but most of them only consider a single dergee of freedom (sdof) main structure. the first design method for stifness of tmd spring was proposed by den hartog [1]. it comes from observation of two invariant points on frequency response function (frf), which are independent of tmd damping. the optimal damping was derived by brock [2] for both motion and force excitation. these settings of parameters are close to the optimal ones for undamped main structure. thanks to simple application, it is still a recomended design method in several modern guidelines [3–5]. the closed form solution of optimal tuning and damping for undamped main sdof system was found by o. nishihara and t. asami [6]. abubakar and farid presented a numericaly derived formula for design of tmd on clasically damped sdof structure [7]. simplification of main structure by sdof model is sometimes insufficient and may lead to wrong tuning or unnecessary big mass of tmd. especially, large structures such as bridges or high buildings, which may be sensitive to dynamic load, cannot be modeled as sdof structures. rana and soong showed the importance of numerical tuning in case of multiple degrees of freedom (mdof) structures and shows the example of multiple tmds, each tuned for particular mode [8]. ozer and royston extended hartog’s idea to mdof systems [9], but the proposed method requires inversion of the dynamic stiffness matrix and numerical solution of system of equations, which may be impractical. this paper presents numerical optimization of single and multiple tmds, on a mdof damped stucture. the goal is to minimize the maximal amplitude (also called h∞ optimization) of the structure caused by simple harmonic load. in the case of h∞ numerical optimization, the problem is that large amount of solutions needs to be evaluated for various tmd parameter values and forcing frequencies. to reduce the camputational complexity of the problem, method proposed in [10] was applied. this method operates with transformation to modal coordinates and allows us to find steady-state response of structures with tmd without solving a system of equations. this method is also suitable for design of tmds for existing structures where excessive vibrations were measured and modal characteristics are known. 2. response calculation the main structure is represented by classically damped mdof system with n degrees of freedom and it is loaded by general dynamic load f(t). the system is described by equation of motion kx + cẋ + mẍ = f(t) , (1) where k, c and m represent stiffness matrix, damping matrix, and mass matrix, respectively, f (t) is the loading vector and x is vector of displacement. the vectors of velocity and acceleration are represented by symbols ẋ and ẍ. 2.1. single tmd tmd is attached to the m-th degree of freedom. the total system can be represented by system of n + 1 equations, but the equation which corresponds to the 100 https://doi.org/10.14311/app.2020.26.0100 https://ojs.cvut.cz/ojs/index.php/app vol. 26/2020 example of an article with a long title mass of tmd can be separated [9], as can be seen in eqs. (2) and (3). [k + ka]x + [c + ca]ẋ + mẍ −kamxa −c a mẋa = f(t) , (2) kaxa + caẋa + maẍa − [cam] t ẋ− [kam] t x = 0 , (3) where: ka defined by eq. (4), ca defined by eq. (5), kam m-th column of matrix ka, cam m-th column of matrix ca, ka stiffness of tmd, ca damping of tmd, ma mass of tmd, xa displacement of tmd, ẋa velocity of tmd, ẍa acceleration of tmd. matrices ka and ca contain only a single non-zero element on the m×m position: ka =   0 0 · · · 0 0 ... ... ... ... · · · ka(m×m) ... 0 · · · · · · 0   , (4) ca =   0 0 · · · 0 0 ... ... ... ... · · · ca(m×m) ... 0 · · · · · · 0   . (5) assuming the loading force is harmonic with amplitude fa and circular frequency ω, we only need to find a solution in the form of steady-state response. the equations (2) and (3) can be transformed to modal coordinates according to eq. (7). substituting xa expressed from eq. (3) to eq. (2), the following result is obtained: [ ω + iωc′ −ω2i+ + (iωca + ka)(−ω2ma) −ω2ma + iωca + ka φtm,∗φm,∗ ] q = φt fa , (6) x = φq , (7) where ω, φ, and c′ are spectral matrix, modal matrix, and modal damping matrix, respectively. φm,∗ denotes the m-th row of modal matrix and q is the vector of modal coordinates. the modal matrix must be mass orthonormal with mode shapes φ arranged as its columns. the spectral matrix contains squares of natural circular frequencies ωj on its diagonal. for a classically damped structure, c′ is also diagonal containing members 2ξjωj where ξj is the damping ratio of j-th natural frequency. eq. (6) is a system of n linear equations. thanks to its structure, sherman–morrison formula can be used to find the inversion of the matrix on the left-hand side. the solution of separate modal coordinate qj with used substitutions is given by following equations: [ ω2ma(iωca + ka)φm,jst2 (−ω2ma + iωca + ka) −ω2ma(iωca + ka)s1 + + φtj ] fa ω2j −ω2 + 2iξjωωj = qj , (8) s1 = n∑ j=1 φ2m,j ω2j −ω2 + 2iξjωωj , (9) st2 = n∑ j=1 φm,jφtj ω2j −ω2 + 2iξjωωj . (10) as can be seen in eq. (8-10), all mode shapes must be known to find an exact contribution of one particular mode shape to the response. however, mode shapes with frequencies, which are far from forcing frequency have negligible influence on the response. therefore using only several selected mode shapes in eqs. (7-10) can lead to very precise results. the reduction of number of mode shapes is necessary for fast response evaluation. 2.2. multiple tmds in the case of multiple tmds, the evaluation of separate modal coordinates using eq. (8) is no longer possible, but sherman–morrison formula a−1i+1 = [ai + uv t ]−1 = a−1i − a−1i uv t a−1i 1 + vt a−1i u (11) can be repeatedly applied to get a vector of modal coordinates q. each addition of tmd requires one evaluation of eq. (11) with substitutions given by eqs. (1214). a0 = ω + iωc′ −ω2i , (12) u = (iωca + ka)(−ω2ma) −ω2ma + iωca + ka φtm,∗ . (13) vt = φm,∗ (14) for j tmds, the vector of modal coordinates is q = a−1j φ t fa . (15) this method may be inappropriate for large number of tmds because of its high numerical complexity caused by multiplication of full matrices. however, its usage still avoids the solution of the system of 101 jan štěpánek, jiří máca acta polytechnica ctu proceedings equations. in section 4.5, this method is successfully used for the evaluation of structural response with three tmds. as can be seen in eqs. (11-14), the system receptance matrix a−1j is found without necessity of matrix inversion. the only matrix which must be inverted is a0, but it is a simple procedure, because a0 is a diagonal matrix. the receptance matrix can be used for fast evaluation of response to various load vectors. 3. simplified tmd design well known simplified design of tmd for minimization of maximal amplitude, first introduced in [1, 2] by hartog and brock, is used in this study as reference method and is compared to numerical optimization of tmds. this method was chosen for its easy application and the fact it is widely used and recommended. using following notation: mef f effective mass of the main structure, ωj n-th natural frequency of the main structure, ωa = √ ka/ma natural frequency of tmd, µ = ma/mef f mass ratio, β = ωa/ωj frequency ratio, ξa = ca/ca,cr damping ratio of tmd, ca,cr = 2 √ kama critical damping of tmd. the optimal values of parameters β and ξa for force excitation are βopt = 1 1 + µ , (16) ξa,opt = √ 3µ 8(1 + µ) . (17) this method was firstly proposed for sdof main structure, but it can be used for design of tmd which is supposed to reduce vibration of j-th mode shape. the effective mass of the main structure mef f can be found according to the following equation: mef f = 1 φ2m,j = φ′tj mφ ′ j , (18) where φ′j denotes the j-th mode shape normalized with respect to the m-th ordinate where tmd is attached hence the m-th ordinate is equal to one. 4. numerical example this chapter analysis a cable stayed footbridge. the structure is suspected of being sensitive to dynamic pedestrian load. a numerical model of the structure was created in matlab. design of tmds was performed both using simplified method indicated in section 3 and numerical optimization. mode frequency [hz] damping ratio [-] 1 0.733 0.0190 2 0.923 0.0148 3 1.300 0.0106 4 1.595 0.0087 5 1.818 0.0078 6 1.924 0.0075 7 2.002 0.0071 8 2.218 0.0063 table 1. natural frequencies and damping ratios. 4.1. model of the structure the structure is 243 m long, symmetric, cablestayed bridge with prestressed concrete bridge deck. a scheme of the structure and numbering of important nodes can be seen in figure 1. two-dimensional finite element (fem) model was created in order to examine vertical vibrations. the model contains nodes which define geometry of the structure and also uses nodes added by fem mesh. the final model contains 810 degrees of freedom. eight lowest natural frequencies, mode shapes and associated damping ratios were measured on the real structure. only vertical bending mode shapes are taken into account in this paper. the numerical model was identified to be in accordance with measurement. natural frequencies and damping ratios are summarized in table 1. 4.2. load as can be seen in table 1, several frequencies are close to 2 hz, which is typical frequency of human walk and thus resonance effect may occure. according to guidebook [5], a stationary load model which represents a group of 8-15 walkers can be used to evaluate the response and to decide whether the vibration is excessive. the group of walkers is simplified to one harmonic force in resonance with mode, which has the frequency closest to 2 hz. the force should be applied to the most adverse position, which may be understood as a place with the highest vertical ordinate of the forced mode shape. in the case given, it was decided to find maximal displacement over the bridge span for three positions, which corresponds to the maximal ordinate of mode shapes. they are node 19 for the 5th mode shape, node 5 for the 6th, and node 27 for the 7th one. the results can be seen in figure 1. the highest peak in range between 1.5-2.5 hz is caused by force situated in node 5 in resonance with 6th natural frequency. therefore the harmonic force with amplitude fv is placed in node 5. fv = kv × 180 = 3 × 180 = 540 n . (19) however, the amplitude fv is important only for taking the absolute value of the response, but it has no impact on the design of tmd. 102 vol. 26/2020 example of an article with a long title figure 1. static scheme of the bridge. figure 2. frfs of force in nodes 5,19, and 27. figure 3. mode shape no. 5. figure 4. mode shape no. 6. figure 5. mode shape no. 7. node 5 is appropriate place for simple harmonic load, because mode shapes 5,6 and 8 also have high ordinate in this node, therefore multiple modes are forced. for simplified tmd design from section 3, the load does not play a role, but it is important for numerical optimization and calculation of frf. 4.3. objective function an objective function must be defined for the optimization of tmd. in this paper, it was decided to use h∞ optimization maximal value of displacement amplitude along the bridge deck between 1.7 and 2.2 hz, where the load causes highest response. that means 103 jan štěpánek, jiří máca acta polytechnica ctu proceedings maximal values of vector of displacement x create frf, which creates a continuous frf with discontinuous derivative. highest peak of frf is then the value of the objective function which is minimized. the step of forcing frequency for evaluating the frf was set to 0.001 hz to catch the sharp peaks but with respect to the speed of calculation. the transformation to modal coordinates allows us to use reduced number of mode shapes to speed up the evaluation of response. observation of frf showed that using more than first 12 mode shapes and natural frequencies does not affect the values of peak response in chosen frequency range. therefore it was decided to use only first 12 mode shapes. the reduction of number of mode shapes reduced the size of the problem dramatically (from 810 equations to 12 equations) without noticeable decrease of accuracy. 4.4. single tmd 4.4.1. simplified design the simplified design method presented in section 3 was performed in order to reduce vibration of the sixth mode shape, because the pacing force causes the highest peak on frf in resonance with the sixth natural frequency, as can be seen in figure 2. tmd with one vertical degree of freedom and mass ma = 600 kg is attached to node 5, and its parameters were designed according to section 3, which resulted into values ka = 86300 nm−1, and ca = 772 nsm −1. the mass of 600 kg is approximately only 0.15 % of the total mass of the structure, but it can reduce the vibration significantly. 4.4.2. numerical optimization to optimize tmd numerically, function fmincon implemented in matlab was used. the function uses interior-point algorithm and is designed to solve constrained multivariable optimization problems. the only implemented constraints were ka > 0 and ca > 0. the function was used to find optimal values of ka and ca which minimize the objective function. generally, in the case of single tmd, the objective function can have at most one local minimum for each natural frequency. in our case, the function has only one local minimum for the sixth frequency, because the peak of frf reaches the highest level in resonance with the sixth natural frequency. the initial values of ka and ca define the local minimum to which interior-point algorithm converges, therefore it is advantageous to begin with tmd tuned close to the natural frequency which response is minimized. for parameters of tmd, which are "far from optimal", a gradient of objective function may be very low, therefore convergence problems may occur, or it can take a large number of iterations to find the local minimum. parameters designed according to section 3 are quality starting point for quick convergence. the parameters, which optimize the objective function ka = 84100 nm−1 and ca = 1259 nsm −1 were mode shape node ka [nm−1] ca [nsm−1] ma [kg] 5 19 26000 98 200 6 5 29100 150 200 7 27 31500 129 200 table 2. tmds designed to reduce vibration of the 5th, 6th, and 7th mode shape. found numerically. corresponding frfs for both methods of design can be seen in figure 6. 4.5. multiple tmds as we can see in figure 2, there are three dominant natural frequencies between range 1.7-2.2 hz. therefore three tmds are designed in this section to show that a higher number of tmds can reduce the objective function more effectively than one. it was decided to keep the sum of tmds mass constantly at 600 kg. the constant sum of weight allows for a relevant comparison of one and multiple tmds. 4.5.1. simplified design in order to reduce frf, three tuned mass dampers were designed according to section 3 to reduce vibration in resonance with the 5th, 6th, and 7th natural frequencies. all of them were positioned to the node where the highest ordinate of damped mode is located. their properties are summarized in tab. 2. the mass was equally distributed among tmds because its optimal distribution remains unknown. 4.5.2. numerical optimization the numerical optimization was performed with two basic settings. the first one uses position of tmds used by previous section (nodes 19, 5, and 27) and was expected to provide the best results. the second one places all tmds in node 5. this calculation was proceeded in order to show that dividing the mass among more tmds can perform better than one tmd, if more than one mode participate on the response. theoretically, one tmd is only a subset of multiple tmds with the same sum of mass. therefore, dividing the mass provided additional room for a performance improvement. the objective function was minimized finding ca,i, ka,i, and ma,i for i =1. . . 3. the following constraints were implemented: • ca,i ≥ 0, • ka,i ≥ 0, • ma,i ≥ 0, • ∑3 i=1 ma,i = 600 kg. from the numerical point of view, the problem is much more complicated than the one with single tmd because 9 variables were optimized. moreover, the function contains multiple local minima. therefore, 104 vol. 26/2020 example of an article with a long title settings no. design method node ka[nm−1] ca [nsm−1] ma [kg] maximal displacement [mm] no tmd 3.447 1 simplified 5 86300 772 600 1.349 2 numerical optimization 5 84100 1259 600 1.103 3 simplified 19 26000 98 200 1.5435 29100 150 200 27 31500 129 200 4 numerical optimization 19 26800 213 208 1.1005 24500 198 157 27 33500 312 236 5 numerical optimization 5 29800 191 189 0.9205 21600 147 176 5 34000 391 235 table 3. comparison of design methods and peak response between 1.7-2.2 hz. it was decided to run the optimization multiple times from one hundred randomly generated initial points, which fulfilled the constraints. this procedure allowed us to find good results, but it is not possible to say that the global minimum was found. 4.6. results the results are summarized in table 3. a total of 5 different tmd settings were found using simplified method and numerical optimization. as expected, tmds designed using den hartog’s criteria were able to reduce the vibration though they were originally proposed for structure with one degree of freedom. it also can be noticed in figure 6 that single tmd can positively affect vibration of adjacent mode shapes if their frequencies are close enough, and if the adjacent mode shapes have sufficiently high ordinate in the position of tmd. figure 6. frf settings no. 1-2. as is shown in figures 6 and 7, the numerical optimization was able to improve the performance of both single and multiple tmds. in the case when starting parameters of tmds were randomly chosen, we expected that each tuned mass damper would converge to a point where it is able to damp one of the figure 7. frf settings no. 3-5. dominant frequencies. this assumption was confirmed and can be seen in figure 8, which shows the frfs of tuned mass dampers. figure 8. displacement of tmds settings no. 5. unexpectedly, the best results were provided by settings where all three tmds were placed in node 5, despite the fact that this is not the place where the highest ordinates of some mode shapes are located. this phenomenon can be explained by closer observation of mode shapes in figures 3, 4, and 5. the force is placed in node 5, therefore all the active mode shapes oscillate with the same phase in this 105 jan štěpánek, jiří máca acta polytechnica ctu proceedings node. however, in nodes 19 and 27, the sixth mode shape oscillates in opposite phase than mode shapes 5 and 7. therefore neither node 19 nor node 27 are the best place for tmd. however, this situation is closely connected to the fact that the optimization was performed to reduce vibration caused by only one specific force. for design of this footbridge, symmetric placement of smaller tmds to both sides of the bridge would solve the problem. 5. conclusions modified mode superposition method was used to evaluate response of structure with tuned mass dampers to speed up the response calculation. utilization of this method provided a powerful tool for numerical optimization, which requires numerous evaluations of steady state response. this response evaluation method provides several important advantages in comparison to standard response calculation: • reduction of numerical complexity without noticeable error thanks to reduced number of mode shapes used. • allows for a fast evaluation of various load vectors, because the system receptance matrix is known. • provides possibility to use experimentally measured mode shapes and natural frequencies for more precise response estimate. • can be used for both single and multiple tmds. simplified tmd design method proposed by den hartog was compared to numerical optimization. the results demonstrate that numerical design is more appropriate for mdof structures and can improve the overall performance of tmds. in all cases of comparison, the analysis showed that optimal damping is higher than the one designed by simplified criterion. further it was shown that multiple tmds with the same sum of mass as a single tmd can provide better results. the best settings of tmds was able to reduce the peak response between 1.7-2.2 hz by 73 %, with tmds total mass of only 0.15 % of the structural mass. this result shows that the common recommendations for tmds to have 1-3 % of mass of the structure may lead to uneconomical design. acknowledgements the authors gratefully acknowledge support from the czech technical university in prague, project sgs19/032/ohk1/1t/11 development and application of numerical algorithms for analysis and modeling in mechanics of structures and materials. references [1] j. p. den hartog. mechanical vibrations. mcgraw hill, new york, 1934. [2] j. brock. a note on the damped vibration absorber. journal of applied mechanics 68(a):284, 1946. [3] setra. footbridges: assessment of vibrational behaviour of footbridges under pedestrian loading, 2006. [4] e. caetano, a. cunha, w. hoorpah, j. raoul. footbridge vibration design. crc press, 2009. [5] fib bulletin 32: guidelines for the design of footbridges, 2005. [6] o. nishihara, t. asami. closed-form exact solution to h infinity optimization of dynamic vibration absorber: ii. development of an algebraic approach and its application to a standard problem. proceedings of the spie, 3989, 2000. doi:10.1117/12.384589. [7] i. m. abubakar, b. j. m. farid. generalized den hartog tuned mass damper system for control of vibrations in structures. earthquake resistant engineering structures 104:185–193, 2009. [8] r. rana, t. soong. parametric study and simplified design of tuned mass dampers. engineering structures 20(3):193–204, 1998. [9] m. b. ozer, t. j. royston. extending den hartog’s vibration absorber technique to multi-degree-of-freedom systems. journal of vibration and acoustics 127(4):341–350, 2005. doi:10.1115/1.1924642. [10] j. štěpánek, j. máca. dynamic response of structures with tuned mass dampers in modal coordinates. vibroengineering procedia 23:13–17, 2019. doi:10.21595/vp.2019.20672. 106 https://doi.org/10.1117/12.384589 https://doi.org/10.1115/1.1924642 https://doi.org/10.21595/vp.2019.20672 acta polytechnica ctu proceedings 26:100–106, 2020 1 introduction 2 response calculation 2.1 single tmd 2.2 multiple tmds 3 simplified tmd design 4 numerical example 4.1 model of the structure 4.2 load 4.3 objective function 4.4 single tmd 4.4.1 simplified design 4.4.2 numerical optimization 4.5 multiple tmds 4.5.1 simplified design 4.5.2 numerical optimization 4.6 results 5 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.32.0042 acta polytechnica ctu proceedings 32:42–48, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague assessment of shear strength parameters of soil from load micropile tests karel vojtasík všb – technical univerisity of ostrava, faculty of civil engineering, deparment of geotechnics and underground engineering, ludvíka podéště 1875/17, 708 33 ostrava-poruba, czech republic correspondence: karel.vojtasik@vsb.cz abstract. design and evaluation of the bearing capacity of micropiles are based on shear strength parameters of the soil within which a micropile is installed. these parameters are mostly received from laboratory tests or as an expert evaluation relying on soil description, classification and collected experience. this paper presents an approach that determines cohesion and angle of internal friction from diagrams of displacement on load, these being obtained by compression and traction load tests of real micropiles installed in the known soil. furthermore, the diagrams are interpreted by the mohr-coulomb strength criterion and the desired values of shear strength are thus derived by using it. keywords: displacement on load diagram, load test, micropile, mohr-coulomb, shear strength, soil. 1. introduction bearing capacity of the micropile is determined by the state and shear strength parameters of the ground within which a micropile is installed. the bearing capacity of the micropile is evaluated in terms of limit stability states regarding either peak or residual stability. the achievement of the respective state is conditioned by the displacements of micropile in the ground. calculations of the bearing capacity most often rely on data obtained either on samples tested in the laboratory or on an expert estimate by classification. both these input data do not take into account state conditions completely and indeed has never given technological procedures for micropile installation, yet both of which also contribute significantly to the bearing capacity. a more reliable picture of the micropile bearing capacity is offered by a load test on site, which will give a respective diagram of the displacement on load. unlike computational methods based on laboratory or expert data, this approach encompasses the state conditions and technological procedures of micropile installation. this diagram of the displacement on load comprises also the peak and residual limit stability states. the respective load forces extracted from the diagram may be utilised for further back analysis to evaluate the shear strength parameters of soil. the article here presented sets out a method employing a force calculation formula of micropile bearing capacity and an interpretation of the diagram following the mohr-coulomb criterion of shear strength. 2. introductory comment on method there is a fundamental assumption that the shear strength development in soil ground should be generally the same regardless of whether it is a laboratory test (figure 1) or a practical field situation, such as in the case of a micropile bearing capacity. figure 1. diagram of box shear tests. determination of strength parameters is based on results of two types of tests performed in a certain order on the same micropile. the compression test is performed first, followed by the traction test. this order of testing ensures that during the compression test both strength parameters are actively combined and at the end of this test procedure the cohesion of soil ground is derived completely. in the traction test that follows immediately there can be mobilised only the angle of internal friction because, after the previous pressure test, the cohesion on the shear surface became zero. the output of both kinds of test are the diagrams of displacement on load (figure 2) that have an analogous course with the shear strength on the displacement course of the soil sample obtained by standard 42 https://doi.org/10.14311/app.2022.32.0042 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 32/2022 assessment of shear strength parameters box shear tests. figure 2. graphs of the course of loading tests of a micropile. a second assumption is that the peak limit state contains both soil shear strength parameters and the residual limit state contains only one, namely the angle of internal friction. the peak limit state corresponds to the load of the peak force fp that represents the maximal recorded value during the test. the residual limit state corresponds to the load of the force fr that occurs at a displacement of 25 mm of micropile. a third assumption consists in analysing the peak force fp and the residual force fr by the shear strength morh-coulomb criterion τ = c + σ tanφ. the peak force fp extracted from diagram of displacement on load by compression test there is derived a peak shear resistance tp at the tube shape surface of the micropile and this equals the sum of a shear resistance due to cohesion tc and a shear resistance due to angle of internal friction tφ. tp = tc + tφ (1) the peak shear resistance is given by the peak force according to the formula: tp = fp πdl (2) where: d diameter of tube shape surface of mircopile l length of tube shape surface of micropile (length of micropile) the shear resistance due to the angle of friction is present in both kinds of test during the whole course of tests so the shear resistance due to angle of internal friction is gained by the residual force according to the formula: tφ = fr πdl (3) the resistance due to cohesion only is the difference between the peak shear resistance and the shear resistance due to angle of internal friction. the value of this difference is the cohesion. tc = tp − tφ = c (4) the residual force fr cannot be applied for evaluation of the angle of internal friction φ because the shear resistance due to the angle of friction involves also normal stress σ at the outer surface of the micropile body (tφ = σ tanφ). the normal stress is hard to verify experimentally on site by measurement. also, a common exact determination encounters a problematic coefficient of lateral pressure, for which there is no point in determining in the usual way, because the soil has been grouted with high grouting pressure and the grouting medium has undergone a phase transformation from a liquid to a solid one. thus, the possible way to determine the normal stress and also the unknown angle of internal friction in question offers a variation in approach with variation parameters σ and φ. the key parameter in the variation calculations is the peak force fp as it includes the both cohesion and angle of internal friction. the peak force is put into a modified formula 1. the remaining parameters are either derivable as tc or measured fp. fp πdl = σ tanφ (5) the variation approach provides many combinations of values σ and φ satisfying the formula 5. finally there are selected those ones fitting the relevant values expected at the respective ground site. 3. description of tests the displacement on load diagrams were processed with data obtained experimentally from two series of tests. a group of four micropiles was first subjected to a compression load force and after this the micropiles were loaded again, this time with a traction load force. the micropiles were installed on the site of the external testing frame at všb tuo, fast. 3.1. micropile the tests were conducted on a micropile from a steel hollow rod of a spindle shape casing, type r32. the length of each of the four micropiles was 3 meters. each micropile ended with a drill bit of diameter 45 mm. the micropile was installed by percussion drilling and after subsequently grouted from the micropile tip with a synthetic epoxy resin. the grouting pressure was about of 8-10 mpa and delivery of the resin was terminated when it ascended to surface (figure 3). this design is widely implemented in practice to enhance the bearing capacity of a foundation. 43 karel vojtasík acta polytechnica ctu proceedings figure 3. micropile from steel hollow rod of a spindle shape casing (type r32). atop of the micropile steel rod is incorporated a concrete prismatic cube. the concrete prismatic cube is necessary to facilitate the installation of a hydraulic press on the lender micropile body and to secure a coaxial transfer of compression load force in the micropile. the smallest deviation in coaxial transfer causes a bending of a slender micropile body and it consequently leads to serious distortion in the values of micropile axial displacement. the concrete prismatic cube has dimensions 40x40x50 cm (50 cm is the height of the cap). in the concrete prismatic there is embedded a short section micropile equipped with steel plate and a nut to fix the plate (figure 4). figure 4. concrete prismatic cube. in the traction load force test a traction load force is applied directly on the end of the micropile. the micropiles are installed inside an outdoor steel test frame. the micropiles are positioned in a square layout and the spacing of the micropiles axes is 1.2 m. the distance of the micropiles to the steel test frame foundations is 1.34 m. 3.2. ground all micropiles are installed in a ground profile that consists of an artificial layer of compacted clayey sand of thickness approx. 0.8 m, which replaces the original top layer consisting of anthropogenic sediments and soil with a high organic content (topsoil and sub-soil). the concrete prismatic cube is placed in it. then there follows a monotonous layer of loess to a depth of approx. 5.5 m and glacial sand and gravel are below this. the parameters of the loess are summarised in table 1. 3.3. compression load test the compression load test is carried out on the micropile together with the concrete prismatic cube on the micropile head. a compression load is generated by a hydraulic press, which is propped against the horizontal beam of the steel test frame through a steel support extension. the hydraulic press is placed on a steel plate with dimensions of 0.4 x 0.4 x 0.03 m. its role is to avert possible crushing of the top of a concrete prismatic cube and eliminate deformation when a higher load force sets in. it provides a base for four potentiometer transducers placed at plate corners to measure micropile settlement. the force is read out by a hydraulic press pressure sensor. the installation of the hydraulic press on the micropile is shown in figure 5. figure 5. installation of hydraulic press on the micropile at the micropile compression load test. all four micropiles were tested with the same regime of applying compression forces. the compression forces were increased in steps and after each step there followed a time break which was necessary to accomplish the full settlement of the micropile in each respective force step. the value of a force in the step was the same in all the steps and it was 10 kn. the break time was 5 minutes and it proved to be sufficient to complete the full micropile settlement in the respective force step. the compression test took place until the value of the 25 mm settlement of the micropile was exceeded. additionally, a concrete prismatic cube without a micropile, placed close to the micropiles in the same soil ground, was tested in a compression test. it was tested with the same instrumentation and load regime and during the test a respective diagram displacement on load was recorded. the diagram displacement on load of only the micropile without the concrete prismatic cube is derived as the difference of displacements of the micropile together with the concrete prismatic cube and a concrete prismatic cube. figure 6 shows the diagrams 44 vol. 32/2022 assessment of shear strength parameters h or iz on d ep th [m ] d en si ty ρ [g /c m 3 ] p or os it y n [% ] w at er c on te nt w [% ] li qu id li m it w l [% ] p la st ic li m it w p [% ] sa tu ra ti on s r c oh es io n c e f [k p a] a ng le of fr ic ti on φ e f [◦ ] d ef or m at io n m od ul us e d e f [k p a] 2.0 1.98 40 19.8 31.5 16.5 0.8 8 0 0 2.7 2.07 39 20.7 36.9 16.0 0.87 8 25 6900 table 1. the parameters of ground of displacement on load measured on the micropile together with the concrete prismatic cube (fm+h), on the sole concrete prismatic cube fh and the derived diagram for micropile only fm. figure 6. diagrams of the compressions tests. 3.4. traction load test the traction load test is carried out on the same micropiles shortly after the compressive load. the concrete prismatic cube is taken away and a traction rig installed above the micropile on ground around the micropile. the traction rig consists of a steel trestle specially constructed for this purpose. the trestle bears a hydraulic press and on its piston there is attached a traction mechanism assembled from two thick steel square plates and four steel rods that join both plates holding stiff together. one plate rests on a press piston the second is fixed with a screw to the micropile. the assembly of the traction rig and its installation on the micropile is shown in figure 7. figure 7. installation of the traction rig at micropile traction load test. the regime of applying traction forces was identical with the compression load test. the only difference between both kinds of tests is that traction displacement of the micropile is measured just with two potentiometer transducers. both transducers are placed on a lower steel plate close to the micropile top in a straight line going approximately through a micropile axe. after traction tests all four micropiles were pulled out from the ground at the end of the traction test. the diameters of the micropile column body along the whole length vary in the range between 70 to 80 mm. the steel rods are coated completely with grouted soil , from the bottom of the concrete prismatic cube to the drill bit. (figure 8). 45 karel vojtasík acta polytechnica ctu proceedings figure 8. the state of the tube shape surface of the micropile after being pulled out of the ground. 4. results the input data assigned straight from the diagram of tests (figure 2) are the load force of the peak bearing capacity fp taken from the diagram of the compression load test. it amounts to 59 kn. another is the load force at the residual bearing capacity fr. it is read as a displacement of micropile of 25 mm. in the case of a compression load test it amounts to 21 kn, while in the case of a traction load test it is greater and amounts to 32 kn. the last straightly assigned parameter is the diameter of a tube shape of micropile determined by measurement after the extraction micropiles out of the ground. this amounts to about 0.075 m (figure 8). the last parameter is the length of the micropile which is 3 m. before performing the calculations according to the relationships ??, it is necessary to explain the different course of the diagrams from both tests and further justify why there was no agreement between the courses and values of residual force in the residual strength stage, as would be expected. the diagram of the traction test does not display any peak in load force in contrast with the diagram for the compression test because during this test cohesion of soil ground should be zero, while it has been drawn out in the preceding compression test. the peak force at the compression test is due to cohesion at the tube shape surface of the micropile and resistance as the ground acts against the micropile base tip. the base is not large and therefore its contribution is disregarded in the analysis. the apparent difference in residual load forces at both tests is probably due to the different development of pore pressure on the tube shape surface of the micropile. in the pressure test the pore pressure will be positive, while in the tensile test it will be negative due to a suction. unfortunately, the pore pressure on the tube shape surface of the micropile is almost impossible to determine by measurement on site. the suction influences and raises the value of the measured residual force in a traction test. this is a reason why the measured value forces at the traction test must be corrected for the suction effect. the suction acts as cohesion on the tube shape surface of the micropile along with friction. the correction force due to suction fs is chosen intentionally as the difference between the residual forces of both tests for the 25 mm displacement and this is about 12 kn. the measured forces in the traction test are all reduced exactly by this value. the cohesion is calculated according to the modified formula 4 complemented with the effect of suction c = fp − (frt − fs) πdl (6) where frtis the residual force of traction test the cohesion is calculated from the values read from the graph of forces at a displacement of 8 mm at which the loading force in the pressure test reached a maximum. c = 59 − (33 − 12) π × 0.075 × 3.0 = 58kpa (7) the variation calculations formula is 59 π × 0.075 × 3.0 − 58 = σ tanφ (8) the result of variational calculations that satisfy the condition given by formula 8 is presented by the graph in figure 9. the relevant combinations of discovered parameters can be read from the graph. figure 9. the graph of combinations of values σ and φ satisfying formula 8. the determination of an angle of internal friction depends on the choice of a magnitude of the normal stress on the tube shape surface of the micropile. 46 vol. 32/2022 assessment of shear strength parameters in the case of a standard situation this magnitude changes with depth and should be same as the horizontal earth stress at rest. the micropile is grouted with a grouting pressure of 8-10 mpa, which is several times higher than a horizontal compound of earth stress value at geostatic stress state in the ground. further, the type of ground has 40% porosity and a degree of saturation between 80 and 90%. thus, there might be expected a normal stress definitely greater than horizontal earth stress at rest. keeping these facts in mind the normal stress might achieve a state where the horizontal compound of stress might be the same as the vertical compound of stress. an effect of grouting might be that the horizontal compound of stress does not depend on the depth and might have approximately constant magnitude along the micropile. the values on the axis σ range from 30 to 60 kpa, and these are consistent with the vertical compound of stress at a depth in the middle of the micropile and respectively at the micropile tip. a photograph on the figure 8, displaying the micropile after being pulled out of the ground and that was used to determine the diameter of it, can be also helpful. the tube shape surface of the micropile is visibly not smooth but rough. the roughness of the tube shape surface of the micropile can contribute to a decision about the magnitude of angle of internal friction indirectly. for the normal stress of 45 kpa, which is a middle point within the range 30 and 60 kpa, the respective angle of internal friction is 28°. this is slightly greater by about 3°than the value of 25°established at the laboratory which complies with a normal stress of 50 kpa. 5. discussion the derivation of the diagram of displacement on load for compression load of the micropile is based on an assumption that the total work performed by the micropile together with the concrete prismatic cube wt must be the sum of the work of the concrete prismatic head wh and the micropile without concrete prismatic cube wm. wt = wh + wm = fm+hs = fhs + fms (9) where: s displacement fm+hcompression force at the micropile together with the concrete prismatic cube fhcompression force at concrete prismatic cube only fmderived compression force at micropile only without concrete prismatic cube fm = fm+h − fh (10) the laboratory data (table 1) do not correspond exactly to the results derived by the presented method. there is an obvious big difference in cohesion. the value of cohesion received from the laboratory test of 8 kpa corresponds with soft consistency and the value seems to be low with regard to a stiff consistency state of soil ground on the testing site. the laboratory value is likely devalued by sampling, transport, and handling of the ground sample before testing. another reason for the different values may be that the force boundary conditions in the laboratory test did not correspond to those on the site. the cohesion value of 58 kpa derived from the compression test is significantly greater than the laboratory value but not outside the range of the stiff soil ground consistency. in the value of the angle of internal friction there is a satisfactory agreement with the range of possible values of normal stresses and also the value determined in the laboratory. 6. conclusion the paper gives us knowledge that the search for, and finding a solution to, a specific geotechnical situation lies in the suitable application of general basic principles. in our case, the derivation of the shear strength parameters of soil ground was based on analysis of diagrams of displacement on load of micropiles. the parameters were derived from comparison with appropriate diagrams between shear strength and displacement of consolidated and soft soil that result from box shear tests and interpretation of the respective diagrams by the mohr-coulomb strength criterion. the results of shear strength determined according to the method were compared with laboratory results of soil samples taken at the testing site. the agreement between the two parameter determination approaches was only not found for cohesion where there occurred an apparent discordance. this discordance in the values of cohesion points to the low reliability of the determination of cohesion in the laboratory. this is due to boundary conditions under which the test is performed and the quality of soil sample. boundary conditions should be the same as on the actual construction site and the quality of the soil sample should be intact. both are difficult to achieve at the same time in laboratory tests, because the conditions on the construction site are difficult to reproduce in their entirety in the laboratory and the quality of the sample is also affected by the handling of samples. all this is a major obstacle to finding a reliable value of cohesion. if the cohesion assessment is based on site tests, both the boundary condition and the quality of the sample are implicitly included in the test. execution of compression and traction tests on the same micropile at the construction site and evaluation of their diagrams displacement on load open up a simple and fast way to a quite reliable determination of shear strength parameters of soil ground and surmount a number of important factors that affect 47 karel vojtasík acta polytechnica ctu proceedings these parameters without being excluded from the evaluation or else having been determined to be of low reliability. this method brings a different approach to determination of the parameters in question or also their verification if they were determined by other methods. acknowledgements the paper has been funded with the support of conceptual development of science, research and innovation in 2021, assigned to the všb-technical university of ostrava, the ministry of education, youth and sports in the czech republic. 48 acta polytechnica ctu proceedings 32:42–48, 2022 1 introduction 2 introductory comment on method 3 description of tests 3.1 micropile 3.2 ground 3.3 compression load test 3.4 traction load test 4 results 5 discussion 6 conclusion acknowledgements acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0007 acta polytechnica ctu proceedings 26:7–12, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app on tuning the dynamic load balancing fem framework michal bošanský∗, bořek patzák czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: michal.bosansky@fsv.cvut.cz abstract. developments in computer hardware are currently bringing new opportunities for numerical modelling. the current trend in technology is parallel processing making use of multiple processing units simultaneously to solve a given problem. this paper deals with exploring the parallel dynamic load balancing framework implemented in the finite element software. this framework is based on a domain decomposition paradigm for distributed memory model. the paper describes the improved technique to determine the actual processor weights related to performance of individual processing units. the load recovery consisting in mesh (re)partitioning is based on actual processor weights. the (re)partitioning process has to be performed during the simulation and whenever the load imbalance is significant. the performance of the proposed technique is tested on the benchmark problem and discussed. keywords: distributed memory, domain decomposition, load balancing, message passing, processor. 1. introduction the main goal of parallel algorithms is partitioning the problem into a set of smaller tasks that can be solved simultaneously. the problem can be partitioned only once before the solution process, which represents static load balancing. when the problem is partitioned during the solution, it represents dynamic load balancing, see [1]. ideal scalability is difficult to obtain because of the overhead cost of the parallel algorithm (synchronization and communication) and because some parts of the problem are essentially sequential. an effective load balancing process that (re)distributes the work to individual computing units is needed in many cases to obtain reasonable scalability. one of the important characteristics of the parallel algorithm is its computational scalability, which is the most important goal in parallel computing. the scalable parallel algorithm achieves a reduction in execution time by using more processing units, ideally in a linear trend. ideal scalability is difficult to obtain because of the overhead costs of the parallel algorithm (synchronization and communication). almost every parallel algorithm has an overhead cost compared to a sequential version. individual tasks cannot be executed concurrently without synchronizing and communicating with other tasks. some parts of the algorithm are also essentially serial and can only be executed by a single thread. in addition to speedup, parallel computing allows large and complex problems to be solved that could not be solved on a single, well-equipped machine. the finite element method (fem) has become a widely used tool for solving problems described by partial differential equations and been widely adopted by engineering and scientific communities as a reliable numerical tool. in mathematics, the fem is a numerical method for finding a solution to boundary value problems for an ordinary differential equation (ode) or a partial differential equation (pde). in the fem, differential equations are converted into an algebraic system of equations by using variational methods aided by decomposition of the problem domain into sub-domains called elements with an appropriate choice of interpolation functions. in many cases the resulting system of equations is nonlinear. in structural mechanics, the non-linearity can originate from nonlinear geometrical relationships (large deformations), from constitutive relationships, and from non-linear boundary condition (e.g. follower type of loading). nonlinear problems are solved incrementally, typically using the n ewton − raphson algorithm. this makes the nonlinear problem solution more demanding than linear problems. the problem is solved in a series of load or displacement increments in which the equilibrium state is iteratively searched. the paper examines the solutions of large scale engineering problems by using parallel computation based on a parallel load-balancing framework. this framework can significantly reduce computational time by using available hardware more efficiently. this paper presents improved algorithm for determining processor weights in load-balancing fem framework. when solving the real problem on the parallel computer, the workload of individual processors can change during the solution. the first source of imbalance originates from the character of the problem. for example, in nonlinear problems, the transition from initial elastic material response to nonlinear regime is often associated with increased computational cost and this 7 http://dx.doi.org/10.14311/app.2020.26.0007 http://ojs.cvut.cz/ojs/index.php/app michal bošanský, bořek patzák acta polytechnica ctu proceedings transition is often associated only to certain regions of overall domain. the second source of load imbalance includes external factors, which can change performance of individual processing nodes or communication network. this typically happens in non-dedicated cluster environments, where processing nodes and communication infrastructure is shared between users. in both cases, the gradual grow of imbalance can have significant effect on performance and on scalability. the load-balancing process is driven not only by emerging applications but also by emerging parallel architectures. these parallel architectures span many scales. for example, clusters have become viable alternatives to tightly coupled parallel computers in small scale systems. on the medium scale, supercomputers are constructed as networks of shared memory multiprocessors (smps) and have complex and nonhomogeneous interconnection topologies [2]. finally, on the largest scale, grid technologies have enabled computations on widely distributed systems, combining distributed clusters and supercomputers into a single computational resource. these grids are a source of extreme computational power and high network heterogeneity. in order to effectively distribute data from any program on such systems, partitioning must account for heterogeneity in the solution environment. load balancing requires evaluation of computing environment parameters (e.g. computing speed, memory and network availability) and determining how to apply this information in the load-balancing process (e.g. adjusting computational domain sizes, selecting partitioning algorithms). the capabilities and performance of the loadbalancing framework depends on many aspects, including problem at hand and hardware configuration and are illustrated on the solution of selected engineering problem. the advantages of implementing this approach are discussed in this paper. 2. introduction and overall design of load balancing framework the models are often solved on complex domains leading to many degrees of freedom and often nonlinear effects have to be taken into account. parallel algorithms should account for load imbalance between particular sub-domains. the imbalance can occur due to several reasons: (i) first load imbalance factor is part of the character of the solution process, as if switching from a linear to nonlinear response in some regions, (ii) second load imbalance factor is an external factor concerning resource relocation, typical in cases when a solution process is run in cluster environments where the available individual computing units are shared by other processes owned by the system or users and lead to changes in allocating processing power. the imbalance is detected in the solution process by monitoring processes during the run time of the process. the decision depends on the amount of load imbalance and the cost of load redistribution. redistributing the work leads to serializing the problem data (elements, nodes, boundary conditions) into messages sent over the network and a receiving process based on unpacking, followed by a topology update reflecting the new work distribution. the idea of parallelization strategy is based on the domain decomposition paradigm. in general, two dual partitioning techniques for the parallel distribution of finite element codes exist, node-cut and element-cut strategy [3]. in this paper, the node-cut strategy, dividing cut dividing the problem mesh into partitions is given in the node-cut technique. the node-cut strategy is assumed when cut dividing the problem domain runs through the nodes. nodes on mutual partition boundaries are called shared nodes, and nodes inside individual partitions are called local nodes. the node-cut partitioning scheme can be interpreted as mesh decomposition using cuts passing through shared nodes of the mesh without crossing any element. the cut strategy ensures duplication of the shared nodes. each processing node is responsible for assembling its contribution to the global system matrix by summing the contributions from individual elements. in order to minimize communication, it is natural that each processing node will assemble and maintain in its local memory the block of global stiffness matrix rows corresponding to unknowns on its partition. this is uniquely defined for local nodes, which are exclusively shared by local elements on individual partitions. for shared nodes, which are shared by elements from multiple partitions, ownership must be defined. in this work, the partition with the lowest rank sharing the shared node is the owner of the shared node and thus responsible for maintaining the corresponding row entry of the global stiffness matrix. it is clear that for shared nodes, the contributions to the corresponding row have to be received from partitions sharing a particular node. the assembly process requires global, unique numbering of equations to be established across the computing nodes in addition to local numbering on individual partitions [4]. partitioning based on the load balancing process can be affected by many factors. ideally, no processor during the parallel solving process should be waiting for another processor to complete the solution. in the typical fem code, the most of the work is proportional to individual elements. the total computational work is determined as the sum of the contribution of individual elements. the computational work of individual element is expressed relatively to computational work of reference element. the load balancing process is based on distributing the overall computational work to a group of available processors with respect to their relative performance. another significant factor concerns minimizing and reducing communication between computational partitions. the assembly of equilibrium equations for shared nodes requires to 8 vol. 26/2020 on tuning the dynamic load balancing fem framework sum up contributions from local as well as remote partitions. the cost of accessing remote memory is much higher than the cost of accessing the local memory. it is therefore important to reduce communication between computational partitions. the present contribution examines the design of an object-oriented framework for dynamic load balancing implemented in oofem code [5]. 3. design of micro benchmarks tests overall processor performance depends on many factors, notably on its frequency, the performance of memory subsystem, and type of code executed. the performance of two cpus can be different for integer and floating point operations, on some smps, some resources are shared between processing cores, etc. the adopted approach to evaluate the individual processor performances is based on a set of so-called micro benchmark tests, that evaluate processor performance for different typical tasks and computing overall performance as the weighted average of performances of individual micro-benchmarks. the processor weights are updated each time, when rebalancing should occur. it is therefore important, that the individual benchmarks have very low time demands and on the other hand provide reliable estimation of individual processors. the first micro benchmark test proposed is a function to compute a numerical integral. in the test, the integral of cosine function is evaluated using numerical integration using newton-cortes formula. the parameter ni defines the number of integration points used to divide the area. therefore, with increasing parameter ni (increasing number of integration points), we have a solution process closely approximating the integral value. this test measures performance in floating point operations. the next type of micro benchmark problem was based on solving a linear system of equations. a fem model of the cantilever beam, divided into neq elements is used to define a linear system. the right hand side of the force vector in our case is a simple vector with its first member set to 1, while all the other elements of the vector are set to zero. the number of equations of the linear system directly depends on parameter neq , which also represents the dimension of the stiffness matrix and right hand side of the force vector. another micro benchmark problem called whetstones is a test that attempts to measure the speed and efficiency of a computer performing floating-point operations. the whetstone benchmark was the first designed for benchmarking [6]. this benchmark is very simple, comprising several sub-tests with active loops executed via procedure calls. this module represents a mix of operations typically performed in scientific applications. the test involves integer arithmetic, floating point arithmetic, "if" statements, calls and so on. at the end of this benchmark, a statement with the results is printed. weights were attached to the different modules (realized as loop bounds for loops around the individual modules statements). the weight distribution of whetstone instructions for the benchmark matched the distribution seen in the program sample. the execution frequency of each module was proportional to the input value of such that the scaling factors for nl = 10 gave the modules a total weight corresponding to one million whetstone instructions for more details see [7]. the final load balancing processor weighs are obtained as a weighted average of individual weights from micro bench-mark tests. the final weights are assembled in the ratio of 15% micro benchmark problem based on a function to compute numerical integral, 80% benchmark problem based on the solution of a linear system of equations and 5% benchmark problem based on measuring the speed and efficiency of a computer performing floating-point operations. the time requirements of individual micro benchmark problems were controlled using parameter n as follows: benchmark to compute numerical integral with parameter ni = 106, solution of a linear system of equations with parameter neq = 103 and computer performing floating-point operations with parameter nl = 6 ∗ 104. 4. dynamic load balancing framework the software design process is an important part of dynamic load-balancing research. the design of a dynamic load balancing framework based on a problem partitioned during the solution process is discussed in this section. dynamic load balancing is typically used for problems involving multiple solution or time steps, where work redistribution is performed to reflect potential work imbalance. partitioning the problem during solution is based on monitoring computation time of individual processors during the solution step. once the solution step is finished, the load balance process is activated by evaluating the measured imbalance. if the imbalance is larger than the user defined tolerance, the (re)partitioning is performed, taking into account the actual performance of processing nodes and processing cost of individual elements. as already noted, when solving a real problem on a parallel computer, the load balance can change during the solution. the only way of reflecting the growing imbalance is to adaptively redistribute work between processing units in order to restore load balance and thereby ensure optimal use of resources. the monitoring and evaluation of imbalance is the task of load balance monitor. this monitor can detect load imbalance by monitoring the time required to perform allocated work on individual processing units. the differences in processing time indicate an imbalance. after imbalance is detected, the decision of whether to restore the load balance or continue is made. this can be 9 michal bošanský, bořek patzák acta polytechnica ctu proceedings a complex task, as load redistribution may in fact be a very complex problem with non-negligible time requirements. the cost of load (re)balancing may be higher than the cost of continuing with a slight loadimbalance. all these aspects have to be considered and are, unfortunately, specific to the problem and implementation. when the imbalance is detected, the work (expressed in terms of individual elements) has to be redistributed to reflect the performance of individual processing nodes. individual elements have assigned weights that reflect the relative computational costs. these depend on element type, material state (elastic, plastic, etc.). these weights are typically obtained in advance from benchmark measurements. the processing power of individual processing units also has to be determined. the load (re)balancing process in this chapter is based on (re)distributing the computational work proportionally to the performance of individual processing units. in principle, additional factors, including available communication bandwidth between individual processing units or available memory, can be taken into account. any load balancing should not only distribute the work according to processing powers but also attempt to minimize the communication cost between individual processing units. in the fem, this means the cuts between partitions (number of shared nodes) should be minimized. failing to meet the secondary criteria can significantly impact overall performance, as the cost of communication (in terms of the time required) is much higher than the cost of computation. load (re)balancing should also attempt to minimize the reallocation of elements as much as possible in order to minimize communication costs. the parmetis library [8] was used in this work, however, the parmetis implementation is not restricted to one specific partitioning library. a parallel partitioner can take advantage of the increased memory capacity of parallel machines (distributed memory model) and improve overall performance. a general load balancing algorithm is responsible for dynamic repartitioning using the processor weights provided by the load balance monitor during the micro benchmark problem and should provide the new partitions with numbers for all local elements on each partition. after the updated element partition assignment is determined, the distribution and classification of nodes also have to be determined. each node is classified as either a local node that remains local on an existing partition or a shared node that is assigned to a remote partition. node classification can be determined from element partitioning. the dynamic load balancing framework implemented in oofem is using the weights based on measuring of individual time involved in the computation. these times represent how long each cpu work on the predefined number of solution steps. the repartitioning is based on default weights set as the ratio of total solution time to individual computational thread time. 5. example using a dynamic load balancing framework the above-mentioned strategy is illustrated on nonlinear 3d finite element analysis of an anchor pullout test. the fe model involved nonlinear nonlocal anisotropic damage model to describe fracture process and consists of 1456 nodes and 16772 tetrahedral elements, which was subsequently refined in 20 steps into a final mesh with 125400 elements and 22441 nodes. the number of solution steps was set to 80 steps. due to the character of the solution, the number of required equilibrium iterations was increasing with progressing solution steps. the number of required equlibrium iterations gradually increased form 9 (first solution step) to 263 (solution step number 80). the existing and proposed processor weight evaluation methods were tested on a workstation (running ubuntu 16.04 os) with an intel ® core ™ i7-4790 cpu @ 3.60 ghz with four cores consisting of two logical processors per core connected to a workstation with an intel® core™ i3-2370m cpu @ 2.40 ghz with two cores consisting of two logical cores. each workstation could simultaneously run a maximum of eight threads and four threads on their cpus and had 15 gb and 8 gb of system memory, respectively. the iterative linear equation solver from the petsc library was used with a block jacobi preconditioner. the dynamic load balancing framework parameters were set up. load (re)balancing was activated after each fifth solution step. the relative wall clock imbalance parameter represents the relative imbalance between wall clock solution time of individual computational threads. when greater than the provided threshold the rebalancing procedure is activated. the additional parameter called absolute wall clock imbalance parameter allows to triggered rebalancing procedure when an achieved absolute imbalance between solution times of individual processing threads is greater than the threshold. the absolute wall clock imbalance parameter was set to 10.0 [s]. the minimum absolute wall clock imbalance parameter allows setting minimum absolute wall clock imbalance threshold for performing rebalancing and was set to 0.5 [s]. the obtained results are presented in tables 1, 2 and 3. in the presented tables, the number of processing units used in the parallel computations is marked as "np". the results obtained using the existing dynamic load balancing framework without the micro benchmark tests to set up processor weights were compared to the results of the computation process without load balancing. these load balancing strategies were finally compared to the presented method based on evaluating the microbenchmark tests. tables 1 and 2 show examples of computation done by processors with hyper-threading [9] technology disabled (intel i3 max. 2 threads, intel i7 max. 4 threads). the computational process with hyper-threading technology enabled on the i3 processor is shown in table 3. 10 vol. 26/2020 on tuning the dynamic load balancing fem framework np no dlb [s] dlb [s] dlb w. [s] 2 1659 1473 1411 3 1408 1399 1230 4 1372 1215 1181 5 1277 1179 1131 table 1. execution times comparing solutions with estimated processing weights (dlb w.), uniform processing weights (dlb) and without a load balancing (no dlb) framework (intel i3 1 computing unit + intel i7 4 computing units). np no dlb [s] dlb [s] dlb w. [s] 4 1524 1479 1373 5 1365 1314 1154 6 1330 1302 1140 table 2. execution times comparing solutions with estimated processing weights (dlb w.), uniform processing weights (dlb) and without a load balancing (no dlb) framework (intel i3 2 computing units + intel i7 4 computing units). np no dlb [s] dlb [s] dlb w. [s] 2 1847 1662 1373 3 1606 1339 1303 4 1610 1340 1435 5 1853 1642 1578 table 3. execution times comparing solutions with estimated processing weights (dlb w.), uniform processing weights (dlb) and without a load balancing (no dlb) framework (intel i3 4 computing units (ht) + intel i7 4 computing units). the results confirm that better performance is obtained when appropriate weights are used. in the case with hyper-threading technology enabled, the scalability trend is not ideal in the solution based on six and eight computational threads and is worse, for example, than the solution based on two and four computational threads. the significant decrease in performance can be attributed to the hyper-threading technology of intel processors, which assembles and shares some cpu resources between hyper-threaded cores and only takes place with 4 threads (note that the workstation with intel® core™ i3 had two physical hyper-threaded cores). however, by using dynamic load balancing and appropriate weights (processor performance parameter), we can achieve better performance than in the solving process without load balancing or with the existing dynamic load balancing framework not using micro benchmark tests. 6. conclusions the contribution based on an improved methodology to determine processing weights parameters as a part of the load balancing framework was presented in this paper. the parallelization strategy was based on the dynamic load balancing process using weights that represented the performance of computational units. the performance of the upgraded dynamic load balancing process (processor weights parameters) was compared to the previously implemented dynamic load balancing process. in this work a nonlinear benchmark problem was considered to evaluate the performance of dynamic load balancing framework. the results showed differences in the performance of the upgraded and previously implemented dynamic load balancing process for the considered benchmark. the upgraded dynamic load balancing process had better performance than the previously implemented dynamic load balancing process. future studies will investigate other testing examples (simple linear equation system solutions) to more precisely assess variables as a representation of computational unit performance. it is clearly a benefit to using the appropriate measures of performance of individual computational units performance of computational units in the load balancing process as a necessary part of any finite element parallel code. the performance of the described load balancing framework could potentially be significantly improved by considering communication speeds between processing nodes by using parameters that represent communication speeds through different, heterogeneous networks. the load balancing framework is based on parallel computation using a distributed memory model. this model uses a message passing interface. for example, many distributed systems are now being constructed using a variety of different communication networks, such as ethernet and asynchronous transfer mode (atm). in addition to this hardware heterogeneity, there is a heterogeneity in the types of messages produced by parallel programs. short synchronization messages require low latency. conversely, large data messages require high-bandwidth, though they can tolerate high start-up latency. the different types of networks have different performance characteristics, while the different types of communication messages may have different communication requirements. the performance of parallel computations based on a distributed memory model is typically limited by communication overhead. high-performance networks and the use of multiple heterogeneous networks can help reduce this overhead. further research could focus on designing the load balancing framework’s input parameters, which represent different types of networks providing a data path between the same pair of network nodes. these load balancing parameters (network weight parameters) may allow a solution process to maximize efficient use of different types of networks rather than passively accept the given features of a single network. 11 michal bošanský, bořek patzák acta polytechnica ctu proceedings acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs16/038/ohk1/1t/11 advanced algorithms for numerical modelling in mechanics of structures and materials. the authors acknowledge the support of the op vvv meys funded project cz.02.1.01/0.0/0.0/16_019/0000765 "research center for informatics". references [1] k. schloegel, g. karypis, v. kumar. parallel static and dynamic multi-constraint graph partitioning. concurrency and computation: practice and experience 14(3):219–240, 2002. doi:10.1002/cpe.605. 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[8] v. k. g. karypis. parmetis: parallel graph partitioning and sparse matrix ordering library. department of computer science, university of minnesota„ 1997. [9] d. h. e. a. d. marr, f. binns. hyper-threading technology architecture and microarchitecture. intel technology journal 6(1), 2002. 12 https://doi.org/10.1002/cpe.605 https://doi.org/10.1016/j.apnum.2004.08.028 https://doi.org/10.14311/app.2017.13.0016 https://doi.org/10.1093/comjnl/19.1.43 https://doi.org/10.1145/998395.998396 acta polytechnica ctu proceedings 26:7–12, 2020 1 introduction 2 introduction and overall design of load balancing framework 3 design of micro benchmarks tests 4 dynamic load balancing framework 5 example using a dynamic load balancing framework 6 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0250 acta polytechnica ctu proceedings 33:250–255, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague axial behaviour of strengthened circular hollow reinforced concrete column with cfrp partial confinement ruqayyah ismaila, raizal saifulnaz muhammad rashidb, fariz aswan ahmad zakwana, ∗, hazrina ahmada, farzad hejazib a school of civil engineering, college of engineering, universiti teknologi mara, cawangan pulau pinang, kampus permatang pauh, 13500 permatang pauh, pulau pinang, malaysia b department of civil engineering, faculty of engineering, universiti putra malaysia, 43400 serdang, selangor, malaysia ∗ corresponding author: fariz838@uitm.edu.my abstract. circular hollow reinforced concrete column will experience deterioration due to many causes such as natural disaster, corrosion, low quality and others. therefore, structural strengthening is required since replacing the deteriorated column is costly. carbon fibre reinforced polymer (cfrp) confinement is one of the common method used for column strengthening. however, being a nonbiodegradable material, the usage of cfrp will contribute to environmental issues. therefore, this study investigates the effect of using partial cfrp confinement in a strengthened circular hollow reinforced concrete column. six specimens of 2 m height with 250 mm and 110 mm for outer and inner diameter were tested. the effectiveness of partial cfrp confinement is analyzed by investigating the tested specimens’ axial and transverse displacement and strain behaviour. this study has shown that with cfrp partial confinement the strength is increased up to 44% from the unconfined circular hollow reinforced concrete column. the results of the study showed that, circular hollow reinforced concrete column with cfrp partial confinement able to enhanced the column load carrying capacity. keywords: cfrp, frp, hollow column, partial confinement. 1. introduction concrete is always known as one of the most reliable materials used in the construction industry. it was always acknowledged as a minimum maintenance material where extensive damage can always be repaired and strengthened. however, many other factors could cause the concrete structure to deteriorate during its service years. circular hollow reinforced concrete column (chrcc) is commonly used as a tower, concrete chimney, large bridge columns and piles, and offshore platforms structures. these structures’ cost could be minimized sufficiently, as the section area and the self-weight is reduced. it significantly increases the sectional moment-of-inertia with an enormous depth and concentrated flanges than regular solid sections with similar areas [1]. to date, cfrp is a well-known strengthening material that has elevated the structural strengthening method tremendously. many experimental works have been carried out resulted from more than 80 concrete frp stress-strain models developed. however, most works focused more on frp stress-strain full confinement but not partial confinement. even though partial cfrp confinement can provide too high strengthening enhancement to the concrete, the number of experimental works is still limited. regardless of all benefits of using cfrp fullconfinement, it does come with several limitations. full cfrp confinement experienced sudden failure explosion without early warning since no initial crack or spalling can be seen. furthermore, when improper full cfrp confinement is installed, it could cause air voids and de-bond between the cfrp and the concrete surface [2]. the use of full cfrp confinement will increase the cost unnecessarily and caused a hazard to the environment due to toxic gasses released during the cfrp manufacturing. it is harmful not only to the environment but also to any people directly in contact with this material either during the manufacturing or installation [3]. therefore, partial crp confinement is seen as a reasonable option compared to full cfrp confinement due to its reduced material used, faster and easy installation, and suitability for a structure requiring moderate strength enhancement [4]. however, there is no proper guideline in designing partial cfrp confinement since experimental and theoretical investigations are still inconclusive [4]. even though several researchers have highlighted the advantages of using partial cfrp confinement ([5–7]) its application is still limited compared to full cfrp confinement. it is due to a limited fundamental understanding concerning the development of stressstrain of cfrp partially confined concrete ([8, 9]). 250 https://doi.org/10.14311/app.2022.33.0250 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 axial behaviour of hollow reinforced concrete column parameter nominal value concrete and steel average compressive strength of concrete, fc (mpa) 35.03 yield strength of the longitudinal bars, fy (mpa) 734.3 yield strength of the transverse bars, fys (mpa) 470.0 cfrp composite system type of frp used unidirectional cfrp sheets elastic modulus of cfrp, mpa 226.3 ultimate tensile strength of cfrp, mpa 3193.27 fracture strain 1.1% thickness of each layer, tf 1.0 mm table 1. material properties. not enough experimental work to validate numerical study in developing partial cfrp-confinement stressstrain concrete model is one of the significant factors that are not commonly used today in practice. 2. experimental program testing was carried out on six circular hollow reinforced concrete columns for this research work. all columns were constructed with two series. the first series is referred to the column specimen without cfrp confinement while the second series referred to the specimen with cfrp partial confinement. the column’s height is 2 meter with 250 mm and 114 mm outer and inner diameter using 30 mpa normal concrete strength. the base of each specimen was fixed with 800 × 800 mm square foundation. eight numbers of 12 mm diameter reinforcement bar with 500 mpa strength were used. 6 mm diameter was used for transverse reinforcement. in obtaining strain data to analyze the specimen’s behaviour, 10 mm strain gauges were installed at every 500 mm height of the specimen. 20 mm lvdts were also installed at each 500 mm height to validate the obtained data. details of the specimen are as in figure 1 below. 2.1. material materials used for this research work were all tested to validate with data provided by the supplier. cube testing was carried out for 7, 14 and 28 days after the curing time to verify the concrete strength. all testings were conducted at material laboratory universiti putra malaysia, serdang according to bs en 12390-3-2009. as for steel reinforcement, a tensile test was conducted for three 12 mm diameter steel reinforcement bar. testing was conducted using instron esh at lightweight structural laboratory, universiti teknologi mara, shah alam with all procedure as in bs en 1008:2005. the displacement rate of 0.001 mm/sec is used to control the machine during the testing until the ultimate tensile strength is reached. all materials properties are tabulated in table 2. figure 1. detail of specimen (dimensions in mm). 2.2. specimen preparation six chrcc specimens were prepared at structural laboratory, universiti putra malaysia (upm), malaysia. all specimens were cast using two different stages with the same batch for each stage. the first concrete stage was for the specimen foundation base and the second concrete stage was for the column. after casting, all columns specimens were cured with intermittent water spray every day to ensure hydration process did not affect the concrete strength. detail specimen cast is as figure 1 below. 251 r. ismail, r. s. m. rashid, f. a. a. zakwan et al. acta polytechnica ctu proceedings figure 2. partial cfrp installation; (a) adhesive application on cfrp strip (b) cfrp strip installatin on concrete surface. figure 3. failure behaviour of chrcc before and after concentric loading. 2.3. cfrp installation every column specimen surface was examined and grounded with sandpaper to ensure a clear surface before the cfrp installation is carried out. to ensure there is no dust or unwanted particles before applying primer to the column surface, acetone is used. a layer of primer was first applied to the column specimen. after at least 24 hours the specimen was left to dry, partial cfrp strip with 60 mm width was cut and adhesive mixture (4:1) was then applied on the cfrp strip before being applied on the column surface (figure 2). 2 layers of cfrp sheet was used in this study with the fibres oriented in the hoop direction. the spacing of the cfrp strips is based on study carried out by pham et al. [7]. 2.4. testing set-up all columns specimens were tested at structural laboratory, universiti putra malaysia with consideration of fixe end at the bottom and axial load is applied concentrically on the top part of the chrcc column. testing was carried out with a displacement control configuration with hydraulic jack capacity of 1500 kn. 20 mm lvdts and strain gauges 60 mm, 10 mm and 5 mm were used for concrete, cfrp and steel reinforcement. the location of lvdts and strain gauges, as shown in figure 1. 3. experimental result and discussion 3.1. behaviour of specimens under concentric loading all chrcc specimens were tested until failure. typical failures of all specimens, as shown in figure 3 below. most unconfined column specimens failed due to sudden loss after it was compressed. failure was observed happen at the top of the column. figure 3 shows that unconfined chrcc failed by flexure since the concrete was crushed at the top after the specimens’ internal steel reinforcements yielded. for partial cfrp confinement, the specimen fracture not as severe as the unconfined column specimen. however, specimens with partial cfrp strip is still experienced ruptured at ultimate load due to flexural tension with a delay compared to unconfined specimens. with partial cfrp confinement, initial cracks 252 vol. 33/2022 axial behaviour of hollow reinforced concrete column �e���d�� � � � ��� ��� ��� ��� ��� � � � � � �� �� �� �� �� �� �� /r dg �� n1 � 9huwlfdo�'lvsodfhphqw��pp� /rdg�yv�9huwlfdo�'lvsodfhphqw� �8qfrqilqhg� &+&�����$ &+&�����$ &+&�����$ � ��� ��� ��� ��� ���� ���� � � �� �� �� �� �� �� �� /r dg �� n1 � 9huwlfdo�'lvsodfhphqw��pp� /rdg�yv�9huwlfdo�'lvsodfhphqw��3duwldo� &)53� &)�����$ &)�����$ &)�����$ figure 4. load vs vertical displacement; (a) unconfined chrcc (b) partial cfrp chrcc. � � � � � � � � � � � � � � �d� �e�� � ��� ��� ��� ��� ��� ��� ��� ��� � � � � � /r dg �� n1 � 7udqvyhuvh�'lvsodfhphqw��pp� /rdg�yv�7udqyhuvh�'lvsodfhphqw� �����pp��8qfrqilqhg &+&�����$ &+&�����$ &+&�����$ � ��� ��� ��� ��� ���� ���� � � � � � � � � � � �� /r dg �� n1 � 7udqvyhuvh�'lvsodfhphqw��pp� /rdg�yv�7udqyhuvh�'lvsodfhphqw� �����pp��3duwldo�&)53 &)�����$ &)�����$ &)�����$ figure 5. load vs tranverse displacement; (a) unconfined chrcc (b) partial cfrp chrcc. �e�� � � � � � � � � � � � � � � �d�� ϭ ϱ ϭϭ ϭϱ ϯϭ ϭ͕ϭϭϭ ϭ͕ϭϭϯ ϭ͕ϭϭϰ ϭ͕ϭϭϲ ϭ͕ϭϭθ ϭ͕ϭϭϭ ϭ͕ϭϭϯ ϭ͕ϭϭϰ ^ƚ ƌğ ɛɛ �;d wă ϳ $[ldo�6wudlq��pp�pp� 6wuhvv�yv�$[ldo�6wudlq�������pp��8qfrqilqhg �,�ͳ;ϭϳͳ� �,�ͳ;ϯϳͳ� �,�ͳ;ϯϳͳ� � � �� �� �� �� �� ����� ����� ����� ����� ����� 6 wuh vv ��0 3 d� $[ldo�6wudlq��pp�pp� 6wuhvv�yv�$[ldo�6wudlq�������pp��3duwldo� &)53 &)�����$ &)�����$ &)�����$ figure 6. stress vs axial strain; (a) unconfined chrcc (b) partial cfrp chrcc. specimen ultimateload (kn) vertical displacement (mm) transverse displacement (mm) axial strain (mm/mm) transverse strain (mm/mm) cn-0-1 361.90 12.51 2.75 0.007 0.0017 cn-0-2 340.00 3.48 1.40 0.007 0.0122 cn-0-3 420.10 7.05 0.96 0.012 0.004 cf-0-1 1020.62 19.35 3.64 0.020 0.013 cf-0-2 960.04 21.75 5.74 0.016 0.010 cf-0-3 1029.02 17.69 1.25 0.010 0.007 table 2. summary of test result. 253 r. ismail, r. s. m. rashid, f. a. a. zakwan et al. acta polytechnica ctu proceedings can be monitored in between the cfrp strips spacing. 3.2. load-displacement behaviour vertical and transverse displacement of chrcc were measured along with the height at three critical position for every 500 mm. for each unconfined and cfrp partially confined, three specimens were tested to obtain more reliable data. for vertical displacement for unconfined specimens, maximum vertical displacement was 12.61 mm at ultimate load (361.9 kn) for chc-(1)-a. as shown in figure 4 (a), the result is not consistent among the three specimens. chc-(2)-a and chc-(3)-a behaved more linear compared to chc-(1)-a which leads to lower vertical displacement. for cfrp partial confinement specimens, all three specimens behaved similar which it increases non-linearly up to a maximum load before almost linearly reduced until failed. details value of the maximum load for each specimen with respected vertical displacement are presented in table ??. as for transverse displacement, only lvdt at the mid-height of the column specimen is considered for discussion, as shown in figure 5 below. for unconfined specimens, the behaviour of the three specimens was very different from each other. however, it can be observed, two specimens chc-(1)-a and chc(3)-a experienced a similar trend and recorded transverse displacement of 4.24 mm and 6.53 mm respectively. as for cfrp partially confined, all three specimens behaved differently with maximum transverse displacement experienced by cf-(2)-a with 8.95 mm before it is fractured and failed. however, the other two specimens were only experienced 3.64 mm and 1.21 mm, which is lower than the unconfined concrete transverse displacement. by taking the maximum transverse displacement of cf-(2)-a, it can be concluded that cfrp partial confinement sufficiently strengthened chrcc with an increment of 21% from transverse displacement for unconfined chrcc. 3.3. stress-strain behaviour figure 6 below shows the behaviour of stress-strain for unconfined and cfrp partially confined to chrcc specimens. it can be seen that the axial stress was kept increasing until it reached the yield strain of cfrp strips. for unconfined specimens, the maximum axial strain is recorded 0.012 with other two specimens experienced lower axial strain value. cfrp partially confined, cf-(1)-a shows an increment of axial strain up to 0.0202, which is higher than the other two specimens. the other two specimens experienced slightly lower strength and lower axial strain. it is sufficiently proven with cfrp partially confinement; it can increase the strength and the axial strength of chrcc specimens. as for stress vs transverse strain for chrcc unconfined and partially cfrp confined, the trend is increased non-linearly with hardening behaviour for chc-(2)-a. for unconfined chrcc, the transverse strain has reached 0.0122 at 17.74 mpa while for partially confined it reached 0.013 at 26.25 mpa. at 17.74 mpa for partial cfrp confinement, the transverse strain has only reached 0.0045. this has shown that at the strength applied; lower transverse strain occurred as the cfrp strip sufficiently confined chrcc from further expansion (figure ??). summary of the test results is tabulated in table ??. 4. conclusions this research concludes that with cfrp partial confinement, the strength of chrcc can increase up to 44.4%, which is sufficient for construction industry work. it can also be observed that with three specimens tested for every parameter, each parameter’s behaviour was found different. the maximum value was taken as references for every parameter. with data and result tabulated from table 2, the cfrp partial confinement strength enhancement of chrcc specimen can be achieved with lesser cfrp material used and contributing more to the sustainability of construction material. acknowledgements this work fully sponsored by the universiti putra malaysia (upm) research grant under the putra grant initiative (grant reference no: 9439300). the authors would like to express their most profound appreciation to upm for providing all facilities and human resources during the laboratory work phase. full gratitude dedicated to universiti teknologi mara, cawangan pulau pinang, kampus permatang pauh, pulau pinang. references [1] t.-h. kim, h.-t. kang. seismic performance assessment of hollow circular reinforced concrete bridge columns with confinement steel. journal of the earthquake engineering society of korea 16(1):13-25, 2012. https://doi.org/10.5000/eesk.2012.16.1.013. [2] t. w. park, u. j. na, l. chung, et al. compressive behavior of concrete cylinders confined by narrow strips of cfrp with spacing. composites part b: engineering 39(7-8):1093-103, 2008. https: //doi.org/10.1016/j.compositesb.2008.05.002. [3] t. sen, a. paul. confining concrete with sisal and jute frp as alternatives for cfrp and gfrp. international journal of sustainable built environment 4(2):248-64, 2015. https://doi.org/10.1016/j.ijsbe.2015.04.001. [4] j.-j. zeng, y.-c. guo, w.-y. gao, et al. stress-strain behavior of concrete in circular concrete columns partially wrapped with frp strips. composite structures 200:810-28, 2018. https: //doi.org/10.1016/j.compstruct.2018.05.001. [5] l. p. ye, k. zhang, s. h. zhao, et al. experimental study on seismic strengthening of rc columns with wrapped cfrp sheets. construction and building materials 17(6-7):499-506, 2003. https://doi.org/10.1016/s0950-0618(03)00047-3. 254 https://doi.org/10.5000/eesk.2012.16.1.013 https://doi.org/10.1016/j.compositesb.2008.05.002 https://doi.org/10.1016/j.ijsbe.2015.04.001 https://doi.org/10.1016/j.compstruct.2018.05.001 https://doi.org/10.1016/s0950-0618(03)00047-3 vol. 33/2022 axial behaviour of hollow reinforced concrete column [6] a. m. vasumathi, k. rajkumar, g. ganesh prabhu. compressive behaviour of rc column with fibre reinforced concrete confined by cfrp strips. advances in materials science and engineering 2014:110, 2014. https://doi.org/10.1155/2014/601915. [7] t. m. pham, m. n. s. hadi, j. youssef. optimized frp wrapping schemes for circular concrete columns under axial compression. journal of composites for construction 19(6), 2015. https: //doi.org/10.1061/(asce)cc.1943-5614.0000571 . [8] a. d. mai, m. n. sheikh, m. n. s. hadi. investigation on the behaviour of partial wrapping in comparison with full wrapping of square rc columns under different loading conditions. construction and building materials 168:153-68, 2018. https: //doi.org/10.1016/j.conbuildmat.2018.02.003. [9] h. al-nimry and m. jawarneh, effects of increased slenderness in short heat-damaged rc columns confined with frp composites, mater. struct. constr., vol. 50, no. 1, 2017, doi:10.1617/s11527-016-0965-7. 255 https://doi.org/10.1155/2014/601915 https://doi.org/10.1061/(asce)cc.1943-5614.0000571 https://doi.org/10.1016/j.conbuildmat.2018.02.003 10.1617/s11527-016-0965-7 acta polytechnica doi:10.14311/app.2020.27.0116 acta polytechnica 27(0):116–120, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app electron radiation effect on indentation creep of construction polymers martin ovsík∗, michal staněk, adam dočkal, petr fluxa tomas bata university in zlin, department of production engineering, nám. t. g. masaryka 5555, 760 01 zlín, czech republic ∗ corresponding author: ovsik@utb.cz abstract. cross-linking is a process in which polymer chains are associated through chemical bonds. radiation, which penetrated through specimens and reacted with the cross-linking agent, gradually formed cross-linking (3d net), first in the surface layer and then in the total volume, which resulted in considerable changes in specimen behaviour. creep value is an important parameter that describes the behaviour of a material for the duration of its exposure to long term stress. most of the technical parts used in industry are subjected to the long term stress for their whole life cycle. this can lead to material creep, which directly results in the transformation of dimensions. to eliminate this problem, a number of construction materials was chosen and subsequently irradiated by a source of electrons. this created a 3d network within the polymer structure, which led to an increase of the micro-mechanical and micro-creep properties. evaluation of these modifications was done by state of the art device micro-combi tester manufactured by anton paar. this device lowers the time required to measure the creep by standard technology and it fluently records the changes of indentation depth in relation to applied force. this dependence is then used to calculate the creep values. due to the electron irradiation, a 40 % increase was reached in creep resistance; therefore the useful properties of selected construction materials were improved. keywords: construction polymers, creep, crosslinking, electron radiation, micro-indentation. 1. introduction in the last years, crosslinking of polymers using electron beam radiation has become increasingly used modification method to obtain better mechanical properties. it is an environmentally safe and managed method. during crosslinking, the polymer structure changes, a 3d network is formed from crosslinked bonds and the polymers become more mechanically resistant [1–4]. in the year of 2005, a team of researchers situated in iran investigated the crosslinking of pa6, enriched by 1 to 3 % of crosslinking accelerator triallyl isocyanurate (tac), exposed to accelerated electrons. in this experiment, an electron accelerator with 5 mev was used to expose the samples to radiation doses in range of 40 to 150 kgy. the results showed, that the molecular weight of the polymer samples rose with increasing radiation doses, which was confirmed by the viscosity measurements. furthermore, the gel test indicated, that the pa6 enriched by tac started the crosslinking process even after being exposed to relatively low amounts of radiation. the contents of absorbed water declined with increasing amounts of tac and the absorbed levels of radiation doses [5–9]. maria porubska investigated the radiation effect on pa6 filled with 30% glass fibres and its virgin version. the samples were injection moulded and later irradiated by accelerated electrons with 10 mev energy in doses of 50, 100, 200, 300 and 500 kgy, where 50% of the dosage was applied to each side. the results showed, that regarding the material properties the radiation is more advantageous for unfilled pa6 [10, 11]. the goal of the paper was the study of pa6 irradiated by the electron beams in doses of 0, 66, 99 and 132 of kgy and the radiation effect on micro-creep. 2. methods 2.1. material two types of polyamides were chosen for this experiment, polyamide 6 pa6 v-ptscreamid-11-amn 0 tld and polyamide 66 v-pts-creamid-12-amn 0 tld. these polyamides are used in technical practice. polymer was delivered in granular form, from company pts plastics technology service, germany. to ensure correct meshing, the supplier added 6% of meshing agent (special crosslinking agent taic triallyl isocyanurate) that secures the meshing reaction and creation of 3d network. 2.2. sample preparation process conditions were set according to the manufacturer. these samples were moulded by the injection moulding machine arburg allrounder 470h according to the same process conditions. process conditions (table 1.) all injected materials were dried by dry and transport unit arburg thermolift 100-2. test samples were manufactured with norm čsn en iso 527 in mind. 116 http://dx.doi.org/10.14311/app.2020.27.0116 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 electron radiation effect on indentation creep of construction polymers figure 1. dependence of indentation depth vs. indentation time of tested pa 6. figure 2. dependence of indentation depth vs. indentation time of tested pa 66. parameters unit pa6 pa66 injection pressure mpa 65 80 cooling time s 17 17 mould temperature °c 70 70 zone 1 °c 220 220 zone 2 °c 250 250 zone 3 °c 270 270 zone 4 °c 310 310 table 1. injection conditions. all these samples were sent to the company b. g. s. to germany, where the samples were irradiated. the range of the dosages was set in compliance with experience gained from industrial practice to 33, 66, 99, 132, 165 and 198 kgy. each cycle in the accelerator exposed the test sample to the radiation dose of 33 kgy (3 cycle of 33 kgy = 99 kgy). beta rays are characterized by their ability to irradiate individual packages within a few seconds. all samples were irradiated with electron (beta) rays (electron energy 10 mev). 2.3. micro-indentation creep indentation creep was measured by means of microhardness tester (mht3) by anton paar, switzerland, using vickers indenter tip, according to the csn en iso 14577. on each material at least ten indents were made and the results were statistically treated. standard simple loading-unloading mode was used. the indentation parameters were set according to the standard; see table 2. 117 m. ovsík, m. staněk, a. dočkal, p. fluxa acta polytechnica figure 3. indentation creep. figure 4. gel content. parameters unit value maximum load n 1 load/unload speed n/min 2 holding time s 21600 table 2. indentation parameters. 3. conclusions cit = h2 − h1 h1 · 100, (1) the identification creep was calculated according to the equation 1,where h1 is the indentation depth at time t1 of reaching the test force (which is kept constant), h2 is the indentation depth at time t2 of holding the constant test force [12, 13]. 3.1. gel content a gel content (eq. 2) test is performed in order to determine the non-dissolved gel content of the given material-according to the astm d 2765 standardtest method c. a portion of 0.5 g (of electron-beam irradiated pa 6 and pa 6.6 material) weighed with a precision of five decimal places on a "swiss made ep 125 sm" weighing apparatus (dietikon, switzerland) was mixed with 100 ml of solvent. xylene was used on the pa 6.6 because it dissolves the amorphous part of this material, and the crosslinking part does not dissolve. the mixture was extracted for 24 118 vol. 27/2020 electron radiation effect on indentation creep of construction polymers h. then, the solutes were separated by distillation. after removing the residual xylene, the crosslinking extract was dried for 8 h, in a vacuum, at 100 °c. the dried and cooled residue was weighed again with a precision of five decimal places and compared to the original weight of the portion. the result is stated in percentage as the degree of crosslinking [14]: gi = m3 − m1 m2 − m1 · 100 (2) where gi is the degree of crosslinking of each specimen expressed in percentage, m1 is the weight of the cage and lid in milligrams, m2 is the total of the weights of the original specimen, cage and lid in milligrams, and m3 is the total of the weights of the residue specimen, cage and lid in milligrams. 4. results and discussion the contemporary indentation method allows the measurements of creep behaviour of different types of materials, including the polymers. this method is based on the principle of immediate detection of indentation depth in dependence on time. the differences in the depth, reached by the indentor during pre-set loading force (1n) with holding time 21 600s, can then be used to calculate the indentation creep. figures 1 and 2 display the indentation curves for individual radiation doses for two tested materials (pa6 and pa66). the variance in curves indicates the different behaviour of the tested materials as well as the change of creep behaviour. the results of the indentation creep measurements show (figure 3), that the radiation crosslinking improves the properties of the tested material. the virgin pa6 displayed higher value of indentation creep, while the material exposed to lower amounts of radiation showed a deterioration in properties (66 kgy). on the other hand, the samples exposed to higher amounts of radiation indicated an improvement in said properties, with the test sample irradiated by 165 kgy. the difference between the virgin pa6 and the irradiated pa6 was 16%. finally, the test sample exposed to the highest doses of radiation demonstrated a slight decrease in the creep behavior. for pa66, similar tendencies were found. the worst values of indentation creep were found in virgin pa66. with increasing radiation dosage, the indentation creep was measured to incrementally improve up to the maximum, which was found in test samples irradiated by 165 kgy. the difference between this sample and the virgin material was 39%. samples irradiated by 198 kgy displayed a minor decline in the indentation creep. the results of the indentation test were confirmed by the gel content test (figure 4). the value of gel represents the percentage volume of 3d network created in the structure of tested polymers. the virgin versions of both tested materials contained 0% of gel. on the other hand, for pa6, the maximum value of the gel (79%) was measured in samples irradiated by 165 kgy, which directly corresponds with the results of the indentation test. radiation higher than 165 kgy proved to decrease the content of gel, which was probably caused by the degradation of the material due to the high intensity of the radiation. for pa66, the highest content of gel was measured in test samples irradiated by 99 kgy, which also confirms the results of the indentation creep measurements. the goal of this study is to examine the effect of beta radiation on the indentation creep of the tested polymers (pa6 and pa66). the research that was done proved, that the irradiation of the samples has a positive effect on the creep behaviour of construction polymers which were chosen. the best improvement was measured in test samples irradiated by 165 kgy dosage for both materials. in comparison to virgin material, the indentation creep rose by 16% for the pa6 and by 39% for the pa66. the results of the indentation creep were confirmed by the gel content test, where the highest increase in crosslinking parts was found in test samples irradiated by higher amounts of radiation. the biggest radiation dosage that was used (165 kgy), proved to worsen the indentation creep as well as the gel content, which was probably caused by the degradation of the material due to the high intensity of the radiation. acknowledgements this work was supported by the european regional development fund under the project cebia-tech instrumentation no. cz.1.05/2.1.00/19.0376 and by the ministry of education, youth and sports of the czech republic within the national sustainability program project no. lo1303 (msmt-7778/2014). moreover, it was supported by the internal grant agency of tbu in zlin: no. iga/ft/2020/003. references [1] y. kamran, p.-l. larsson. second-order effects at microindentation of elastic polymers using sharp indenters. materials and design 32(6):3645–3653, 2011. qc 20110520, doi:10.1016/j.matdes.2011.01.043. [2] k. rajeesh, r. gnanamoorthy, r. velmurugan. effect of humidity on the indentation hardness and flexural fatigue behavior of polyamide 6 nanocomposite. materials science and engineering: a 527(12):2826 – 2830, 2010. doi:10.1016/j.msea.2010.01.070. [3] j. schone, d. tondl, r. lach, et al. analysis of pa 6 nanocomposites indentation and creep behavior as a function of temperature and load level using different indentation techniques. polimery 59:722–728, 2014. doi:10.14314/polimery.2014.722. [4] j. g. drobny. radiation technology for polymers. crc press, 2003. [5] o. n. tretinnikov, s. ogata, y. ikada. surface crosslinking of polyethylene by electron beam irradiation in air. polymer 39(24):6115 – 6120, 1998. doi:10.1016/s0032-3861(98)00075-5. 119 https://doi.org/10.1016/j.matdes.2011.01.043 https://doi.org/10.1016/j.msea.2010.01.070 https://doi.org/10.14314/polimery.2014.722 https://doi.org/10.1016/s0032-3861(98)00075-5 m. ovsík, m. staněk, a. dočkal, p. fluxa acta polytechnica [6] g. zamfirova, v. gaydarov, t. zaharescu. microindentation study of electron beam irradiated polyamide samples. chemicke listy 104:283–286, 2010. [7] j. dobransky, p. baron, m. kocisko, et al. solving depressions formed during production of plastic molding. metalurgija -sisak then zagreb54:496–498, 2015. [8] a. j. uddin, y. gotoh, y. ohkoshi, et al. crystal modulus of a new semiaromatic polyamide 9-t. polymer engineering & science 52, 2012. doi:10.1002/pen.22086. [9] k. makuuchi, s. cheng. radiation processing of polymer materials and its industrial applications. wiley, 2011. [10] m. porubska, d. babic, i. janigova, et al. the effect of gamma irradiation in air and inert atmosphere on structure and properties of unfilled or glass fibre-reinforced polyamide 6. polymer bulletin 2015. doi:10.1007/s00289-015-1576-0. [11] m. porubska, i. janigova, k. jomova, i. chodak. the effect of electron beam irradiation on properties of virgin and glass fiber-reinforced polyamide 6. radiation physics and chemistry 102:159 – 166, 2014. doi:10.1016/j.radphyschem.2014.04.037. [12] m. ovsik, d. manas, m. manas, et al. the effect of cross-linking on nano-mechanical properties of polyamide. in polymers and composites in engineering: processing, properties and applications, vol. 699 of key engineering materials, pp. 37–42. trans tech publications ltd, 2016. doi:10.4028/www.scientific.net/kem.699.37. [13] w. oliver, g. pharr. measurement of hardness and elastic modulus by instrumented indentation: advances in understanding and refinements to methodology. journal of materials research 19(1):3–20, 2004. doi:10.1557/jmr.2004.19.1.3. [14] m. ovsik, m. stanek, m. reznicek, l. hylova. study of nano-creep of unfilled and filled cross-linking polypropylene. materials science forum 919:103–110, 2018. doi:10.4028/www.scientific.net/msf.919.103. 120 https://doi.org/10.1002/pen.22086 https://doi.org/10.1007/s00289-015-1576-0 https://doi.org/10.1016/j.radphyschem.2014.04.037 https://doi.org/10.4028/www.scientific.net/kem.699.37 https://doi.org/10.1557/jmr.2004.19.1.3 https://doi.org/10.4028/www.scientific.net/msf.919.103 acta polytechnica 27(0):116–120, 2020 1 introduction 2 methods 2.1 material 2.2 sample preparation 2.3 micro-indentation creep 3 conclusions 3.1 gel content 4 results and discussion acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0160 acta polytechnica 27(0):160–163, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app hot hardness measurements on materials up to 600 ◦c during the first hour of using bruno passilly∗, lara molenda dmas onera, université paris saclay, 29 avenue de la division leclerc f-92322 châtillon france ∗ corresponding author: bruno.passilly@onera.fr abstract. in the aeronautical field, materials are used in severe environmental conditions (temperature, atmosphere, exposure time ...), particularly for engine applications. in order to characterize the use of these materials in the evaluation of their properties, it is necessary to carry out tests in conditions close to their operating environment. hot hardness is a simple method which can be applied on many different materials such as oxidized layers, coatings, composite materials, brazing cords, additive manufacturing materials. onera is developing micromechanical characterization means to carry out vickers microhardness tests from room temperature up to 600 ◦c. in principle, a pyramidal punch is applied on the surface of a material and the applied load is continuously measured during indenter’s moving in the material. the material is tested locally under conditions close to the actual conditions of employment. the goal of this research is to improve microindentation in order to achieve temperature test campaigns up to 600 ◦c under a controlled atmosphere of argon and to validate a method to produce a series of results during the first hour of using up to 600 ◦c. stainless material is studied to compare the evolution of its hot hardness properties versus different parameters such as load, holding time at the maximum load, atmosphere, and thermal duration. a discussion about these measurements and the technical limits of hot hardness technology is presented. keywords: hot hardness, microindentation, stainless steel, temperature. 1. introduction in aeronautic applications, temperatures can reach over 1000 ◦c in work condition. the used materials have to withstand stresses, strains as well as temperature cycles which can lead to the damage of the part. therefore, material characterization plays an important role in the part design. microindentation and hardness tests present many advantages compared to classical characterization methods such as flexural or tensile test: a simple geometry, little material quantity needed, automated data collection and analysis. it can be performed on substrate, thin films or through several layers too [1, 2]. when temperature is increased, it has an impact on the structure and the properties of materials. measuring hardness at high temperature gives important insight of the thermophysical properties of the material. research on micro and nanoindentation at high temperature has increased for the past two decades and many issues remain [3–5]. techniques have been developed to reduce induces thermal dilatation, thermal drift and oxidation which affect the quality of the results. the goal is to validate a method to investigate the changes in hardness versus temperature under the first hour of heating and to avoid thermal drift and oxidation of the indenter or the sample [6–8]. 2. hot hardness set-up 2.1. presentation of the experimental set-up the microindentation machine is composed of the chamber which comprises the sample holder, an optical microscope, the indenter, the indentation station, 2 motorized (x,y) tables to move the sample from the chamber’s opening to the microscope or indentation station, 1 motorized table to move the indenter to the sample and 1 motorized table to move the microscope for focusing (figure 1 and figure 2). a primary pump and a turbo pump are necessary to empty the vacuum chamber. a gas tap is required to introduce argon + hydrogen in the chamber. a water cooling system regulates the temperature of the measurements. a computer and a developed software on @labview are needed to manage all the movements of the tables and to acquire all the datas versus time of the sensors like thermocouples, load sensor, displacement coordinates of the tables. the following table 1 sums up the characteristics of the machine. tmax 800∼◦c fmax 3n atmosphere ar+3%h2 indenter’s shape vickers indenter’s material diamond table 1. characteristics of high temperature microhardness set-up 160 http://dx.doi.org/10.14311/app.2020.27.0160 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 materials hot hardness measurements figure 1. high temperature microhardness synoptic figure 2. high temperature microhardness set-up the two ovens are composed of the heating part, which is made of a platinum coil which is winded around an hollow cylindrical ceramic part. the sample is placed inside the ceramic part of the first oven, and the indenter is placed in the second oven. a current goes through the coil and heated up because of joule’s effect. a thermocouple sensor measures ovens temperature and the controller fixes the working temperature, the slope of the rise ramps temperature and power. as the experiments are carried out at high temperatures, the oxidation of the sample’s surface has to be carefully avoided. in order to work under argon and hydrogen gas, vacuum of the chamber first needs to be done to reach a residual pressure of 5.10−5 mbar minimum. the gas tap is opened and the chamber’s door closed at the same time and the working pressure near from the atmospheric pressure is reached in 3 minutes. the primary pump is turned off once the turbo pump no long turns. 2.2. method the experiments are performed under argon and hydrogen gas at atmospheric pressure. the imprint calibration and the force sensor calibration are made prior to any experiments. the duration of heating must be optimized and controlled for different temperatures. for each oven, the study is performed at 400 ◦c, 500 ◦c and 600 ◦c. the investigated slopes are 15 ◦c/min, 25 ◦c/min and 50 ◦c/min. from each temperature, the time to reach the temperature as well as the overshoot (maximum temperature reached by the oven) is determined. the delay between the oven and the sample’s surface is also determined. all of these temperature measurements are tested on the same material inox 316l-1 (table 2). sample ref. num. temp. test study remarks 316l_1 20, 200,400,500,600 temperature parameters 316l_2 20 applied loadand hardness 316l_3 600 time duration 316l_4 400 hardness 316l_5 500 hardness 316l_6 600 hardness table 2. samples reference 2.3. temperature profile figure 3a shows the temperature profile of the sample’s oven for different ramps and temperatures. it can be seen that the increase in temperature is steady and the set temperature is easily reached: the ovens are well regulated. however, an overshoot is observed. even if it cannot be avoided, it needs to be as small as possible in order not to overheat the sample. to highlight this phenomenon, the curve is zoomed in (figure 3b). the overshoot of the oven is lower than 2 ◦c. therefore, 50 ◦c/min ramp is chosen as settings for all the hardness tests as it allows a faster heating with a reasonable overshoot. 2.4. temperature delay and temperature deviation as the temperature of interest is the temperature of the sample, the following experiment focuses on the temperature delay and the temperature deviation of the sample’s surface to the set temperature. a thermocouple sensor is placed just at the surface of the sample. the sample is placed underneath the indenter and the surface temperature is measured. the investigated temperature is 400 ◦c at 50 ◦c/min. then, the distance between the ovens is decreased to see its influence on the surface temperature. there is 37s seconds delay between the oven’s temperature and the sample’s one (figure 4a). the sample reaches a constant temperature of 400 ◦c with an overshoot which is less than 2 ◦c. however, the temperature difference is 6 ◦c and the temperature does not seem to increase. the temperature deviation between the set temperature and the surface temperature is plotted against the distance between the two ovens (sample and indenter). for all temperatures, the deviation decrease may be observed when the ovens are brought close 161 bruno passilly, lara molenda acta polytechnica figure 3. temperature ramps for 400, 500, 600 ◦c set temperature and 15, 25, 50 ◦c/min ramp (b) overshoot example at 600 ◦c and 50 ◦c/min ramp together. at 2 millimeters the deviation reaches only 1 ◦c difference. for each temperature test, the results can be extrapolated and it can be considered that there is no temperature deviation at contact. 3. hot hardness results on stainless steel sample 3.1. samples preparation the material studied is a stainless steel 316l [9]. the material is cut in a parallelepiped shape 6 mm × 5 mm × 10 mm. sample surface (10 mm × 6 mm side) polishing is performed with polishing papers with decreasing grain sizes from p800 to p4000. diamond slurry is applied on velvet sheets with grain sizes of 3 µm, 1 µm and 1/4 µm. between each step, the sample is cleaned in an ultrasonic bath of water and dried. an optical microscope is used to check the quality of the surface of the sample and to avoid scratches. the last step consists in cleaning the sample in an ultrasonic bath in acetone for 2 minutes and dried. for this study, 5 samples are prepared, sample 1 is used only to develop temperature profile before the study. 3.2. measurement of hardness the following protocol is applied: an imprint is done for a recorded load. each imprint is separated by figure 4. (a) temperature difference between the thermocouple of regulation of the oven and the temperature of the sample (b) temperature deviation as function of the distance between the two ovens 100 µm of the following imprint. the imprints are measured using an electronic caliper square on the printed photo of the screen. for a vickers imprint, each diagonal of the pyramidal imprint is measured separately. hardness is then calculated by using equation (1): hvickers = p avickers = 2.p · sin ψ2 d2 = 1.854. p d2 . (1) with p the load, d the mean of the measured diagonals of the imprint and ψ is equals to 136 ◦. 3.3. study of the effect of the applied load and duration time a series of 5 indents is performed for each following load 50, 100, 150 and 200 grams. equation (1) is used to calculate the hardness. if the mean values are considered, the hardness values are constant as the load is increased for a given temperature of 20 ◦c. the lowest standard deviation is obtained for a load of 150 grams, which will be the load to be applied for all the other tests. the effect of the duration of time at the maximum load is also investigated. a 150 g load and a temperature of 600 ◦c are chosen. the same protocol as the load effect is applied: 5 imprints are done for each load, separated by 100 µm. the results of hardness at 5, 180 and 360 seconds duration time at the 162 vol. 27/2020 materials hot hardness measurements maximum load are constant which show that no effect of duration time at maximum load is occurring. a 5 seconds duration time is chosen for all the following experiments. 3.4. thermal history and changes in hardness the ovens are heated to the set temperature, then the contact point and indentations are made for one hour. the imprints are observed at the optical microscope. the temperature is set to the temperature test, the load used is 150 g and the duration time at maximum load is 5 seconds. for each temperature a new sample of stainless steel 316l is taken to have the same thermal history for each series of points (see table 2). the figure 5 shows an hardness mean of 3.6 gpa at 20 ◦c which decreases at 200 ◦c to reach 2.5 gpa and is stable for a higher temperature up to 600 ◦c. in order to investigate the possible oxidation layer, a series at 20 ◦c was performed after each series at high temperature were done. the values of the second series at 20 ◦c matches the first one. thus, the sample’s oxidation is avoided. figure 5. variation of hardness on stainless steel up to 600 ◦c the final goal is to validate this method to measure the changes in hardness in less than one hour at 600 ◦c. figure 6 represents constant value of hardness over time once 600 ◦c is reached. the mean value of the hardness is 2.4 ± 0.08 gpa. at 20 ◦c there is no more evolution of hardness than at 600 ◦c. 4. conclusions to conclude, the study shows very stable and efficient thermal regulations that allow fast heating and little overshoot which are not reach on conventional microhardness apparatus. oxidation problems are avoided by working under argon atmosphere. changing samples for each temperature is a good method to compare the results with the same thermal history. the tested method seems promising and higher temperatures (up to 1000 ◦c) under ultra-high vacuum or argon atmosphere should be developed to characterize materials which are used under more severe conditions. figure 6. validation of the method at 600 ◦c and 150 g load on a new sample acknowledgements dr k.durst and j.petit for their fruitful conversations and p.lasserre for his technical help. references [1] w. oliver, g. pharr. an improved technique for determining hardness and elastic modulus using load and displacement sensing indentation experiments. journal of materials research 7(6):1564–1583, 1992. doi:10.1557/jmr.1992.1564. [2] a. r. f. jr., g. pintaúde, a. sinatora, et al. the use of a vickers indenter in depth sensing indentation for measuring elastic modulus and vickers hardness. materials research 7(3):483–491, 2004. doi:10.1590/s1516-14392004000300018. [3] j. wheeler, d. armstrong, w. heinz, r. schwaiger. high temperature nanoindentation: the state of the art and future challenges. current opinion in solid state and materials science 19(6):354–366, 2015. doi:10.1016/j.cossms.2015.02.002. [4] t. kutty, k. ravi, c. ganguly. studies on hot hardness of zr and its alloys for nuclear reactors. journal of nuclear materials 265(1-2):91–99, 1999. doi:10.1016/s0022-3115(98)00610-2. [5] r. sundar, t. kutty, d. sastry. hot hardness and creep of fe3al-based alloys. intermetallics 8(4):427–437, 2000. doi:10.1016/s0966-9795(99)00118-1. [6] j. m. wheeler, j. michler. invited article: indenter materials for high temperature nanoindentation. review of scientific instruments 84(10):101301, 2013. doi:10.1063/1.4824710. [7] j. wheeler, r. oliver, t. clyne. afm observation of diamond indenters after oxidation at elevated temperatures. diamond and related materials 19(11):1348–1353, 2010. doi:10.1016/j.diamond.2010.07.004. [8] x. liu. développement de la caractérisation du comportement local à haute température des alliages métalliques par micro indentation. ph.d. thesis, 2017. [9] r. k. desu, h. n. krishnamurthy, a. balu, et al. mechanical properties of austenitic stainless steel 304l and 316l at elevated temperatures. journal of materials research and technology 5(1):13–20, 2016. doi:10.1016/j.jmrt.2015.04.001. 163 https://doi.org/10.1557/jmr.1992.1564 https://doi.org/10.1590/s1516-14392004000300018 https://doi.org/10.1016/j.cossms.2015.02.002 https://doi.org/10.1016/s0022-3115(98)00610-2 https://doi.org/10.1016/s0966-9795(99)00118-1 https://doi.org/10.1063/1.4824710 https://doi.org/10.1016/j.diamond.2010.07.004 https://doi.org/10.1016/j.jmrt.2015.04.001 acta polytechnica 27(0):160–163, 2020 1 introduction 2 hot hardness set-up 2.1 presentation of the experimental set-up 2.2 method 2.3 temperature profile 2.4 temperature delay and temperature deviation 3 hot hardness results on stainless steel sample 3.1 samples preparation 3.2 measurement of hardness 3.3 study of the effect of the applied load and duration time 3.4 thermal history and changes in hardness 4 conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2019.22.0150 acta polytechnica ctu proceedings 22:150–154, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app physical-mechanical properties and durability of hydraulic lime-based mortars with non-traditional biopolymers tomáš žižlavský∗, martin vyšvařil, pavla rovnaníková brno university of technology, faculty of civil engineering, veveří 331/95, 602 00 brno, czech republic ∗ corresponding author: zizlavsky.t@fce.vutbr.cz abstract. the study focuses on the influence of addition of four different biopolymers in various doses (0.1%, 0.5%, and 1%) on the physical-mechanical properties and durability of nhl 3.5 mortars prepared with constant water/binder ratio and the binder/aggregate ratio of 1 : 2 by weight. the flow table test was carried out and the specimens were tested for flexural and compressive strength at the age of 7 and 28 days. the durability coefficient for the flexural strength was calculated after 15 freezing-thawing cycles. all the admixtures have dosage-dependent slump reducing effect. the strengths were increased in comparison with mortar with no admixture, the most significantly in the case of highest dose of carrageenan and diutan gum. all biopolymers with the exception of diutan gum increased the durability coefficient with increasing dosage. sodium salt of alginic acid has been found the most effective in the case of durability improvement. keywords: natural hydraulic lime, sodium alginate, carrageenan, diutan gum, xanthan gum, compressive strength, flexural strength, durability. 1. introduction hydraulic lime has been, as the strongest and most versatile binder, used since the ancient times. natural hydraulic lime is prepared by burning limestone with clay impurities on similar temperatures as a pure limestone. during the burning process the clay decomposes creating silicate and aluminate phases in the binder, which reacts with calcium hydroxide creating minerals similar to clinker minerals in cement. the hydraulic lime hardens firstly by hydraulic reactions and afterwards by carbonation of the rest of the ca(oh)2. viscosity enhancing admixtures (veas) are lately widely studied for the use in concrete to improve properties especially of self-consolidating concrete. their use in lime mortars is far less studied, partially due to the amount of production of concrete in comparison with lime for the building purposes. veas are supposed to improve plasticity of the mortars, their fresh state adhesion, and to eliminate the sagging of renders, mostly by their water-retaining function. [1, 2] the previous results declare, that even though most of the veas are polysaccharides, due to the amount needed to improve concrete properties they do not slow the hydratation reactions of clinker minerals in cement, therefore, if the dosage is similar, they should not affect the hydratation reaction of nhl mortar neither. [1, 3] the current study focuses on the impact of addition of non-traditional biopolymerical veas on the hardened properties and durability of natural hydraulic lime-based mortar. the biopolymers used are microbial based (xanthan gum (xg), and diutan gum (dg)) and seaweed based (carrageenan (cg), and sodium salt of alginic acid (algna)). all of these biopolymers are currently used in food industry as gelling and thickening agents, and stabilisers [4–6]. the most well-known in the building industry are the xg and dg, which are partially in use in self consolidating concrete to reduce segregation and bleeding [4, 5, 7]. dg is more flexible in the use in comparison with xg for its properties are independent on monoand bi-valent (e.g. ca2+) cations concentration as well as the temperature [7–12]. the cationic sensitivity lead to significantly different behaviour of these two biopolymers in lime mortars, where dg improved the strength characteristics and xg notably reduce them, especially the compressive one [13]. the xg was chosen for the research on the hydraulic lime properties, even though it didn’t performed well, mainly because it is reported to have positive influence on concrete behaviour, and the nhl solution contains much less ca2+ cations than the lime putty. sodium and calcium salts of alginic acid have surpassed the commercial super absorbent polymers especially in the way of influence on the hydratation of cementbased mortars and their addition suppressed the same problems as dg and xg addition [8, 14]. in the aerial lime-based mortar [13] algna was one of the best performing admixtures, even though its initial strength reduction with growing dose caused by the highest humidity of the specimens. the use of cg has been reported mainly in geopolymers, for the fly-ash based one its addition increased strengths by creating more condensed structure. [15] this result was lately supported by z. abdollahnejad et al. [16] in the study on alkali activated cement based binder 150 http://dx.doi.org/10.14311/app.2019.22.0150 http://ojs.cvut.cz/ojs/index.php/app vol. 22/2019 physical-mechanical properties of hydraulic lime-based mortars mortars with fly-ash and waste glass addition, where cg addition lead to better workability and also to higher compressive strength, while adding the xg had no positive effect on the strength properties. in the lime mortar cg for the initial strength characteristics didn’t performed well, but showed interesting hint of improvement for the long-term properties. [13] 2. materials and methods 2.1. materials for the preparation of mortar samples same materials as for the previous work [13] were used: the nhl 3.5 lime according to en 459-1 has been mixed with the siliceous sand fraction 0–4 mm (českomoravský štěrk a.s., hulín, czech republic) in a 1:1 volumetric ratio. then the biopolymer was added in the dose of 0.1%, 0.5% and 1% of the weight of binder. the admixtures used were products of sigma-aldrich co (algna) and cp kelco (kelco-crete dg-f – dg, genuvisco cg-131 – cg, and kelzan ap-as – xg). the dry mixture was introduced into water of constant water : binder ratio of 0.7. the flow value was determined according to en 1015-3. the samples were then cast into 40 × 40 × 160 mm prismatic moulds, and demoulded after 48 hours. the beams were stored in the laboratory conditions (20 °c, 50% relative humidity) until the day of testing. 2.2. methods the methodology of the experiments is similar to the previously published results [13]: bulk density of the mortar was determined by weighting and measuring the samples before the strength tests, flexural and compressive strength tests were carried out according to en 1015-11 at 7 and 28 days of age of the mortars. after 28 days of curing time in the laboratory conditions samples underwent the freezing-thawing durability test for 15 cycles according to čsn 72 2452. the reference samples were stored under water to simulate similar conditions as the tested samples. the durability coefficient for flexural strength after 15 cycles has been determined. 3. results and discussion 3.1. flow values addition of any of the admixtures caused workability reduction varying on the dosage and type of the admixture. the results are presented in the figure 1. we can see, that the highest flow value reduction is, as well as for the lime mortar, in the case of dg [13]. the trends are similar to aerial lime-based mortar with exception of xg where the difference is more significant due to the earlier-stated concentration of ca2+ cations in the solution [7, 9–12]. 3.2. bulk density of mortar in the figure 1 there can be seen a decreasing trend in bulk density of the mortar. this is in this case caused by air entrained due to different workability of the mortars, for all of the specimens consisted of the same water : binder : aggregate ratio, the trend corresponds with the flow value reduction one and as has been stated earlier these biopolymers did not show air-entraining function [13]. 3.3. flexural and compressive strength the results obtained by strength tests are presented in figures 2 and 3. the 7 days strength was far less affected, than in the case of lime mortars with different biopolymers [13, 17, 18], but even less affected than the expectations based on the results obtained on cement or nhl mortars found in the literature [14, 19, 20]. all the biopolymers and doses with exception of lowest dosage (0.1%) of polymers for flexural strength surpassed the strength of mortar with none of the biopolymers. for the flexural strength the dg and cg in highest dosage (1%) exceeded significantly all other biopolymers confirming the earlier results on lime mortar [13] and the presumption based on literature research [15, 16]. the xg performed much better then while used in aerial lime-based mortar, but in comparison with other biopolymers, the 28 days results were one of the worst ones. the overall increase of strengths in this experimental set up may be as well interpreted as the impact of dewatering, which can be partially described by the flow value reduction, but the dewatering effect may not cause so enormous difference in strength as in the case of cg and dg [21]. 3.4. durability the results of durability tests carried out after 15 freezing-thawing cycles were summarized in table 1. the durability coefficient for flexural strength was chosen, for it is more representative if the material is used as a plaster. most of the specimens tested survived 15 freezing thawing cycles in the way, they could be tested for flexural and compressive strength. the most damaged were beams prepared from mortar with dg addition of 0.1% and 0.5% (fig. 4) out of the three beams for each mixture only one was compact enough for the three point bending test and even though the test results indicated several damage to the specimens, thus the durability coefficient of dg 1 and dg 5 mortars should be approached with a distance. overall the biopolymer addition lead to improvement of durability as expected based on results found in the literature [3, 17, 18, 22, 23]. the dg addition notably worsens the durability, so it’s even lower than the one of mortar with no admixture. with the growing dosage of biopolymeric admixture the mortar became more durable. the best performing were algna and xg, which even in the lowest dose (0.1%), improved the durability significantly. according to čsn 72 2452 151 t. žižlavský, m. vyšvařil, p. rovnaníková acta polytechnica ctu proceedings figure 1. flow values of fresh mortars and bulk density of 7 days old specimens. figure 2. flexural and compressive strength of mortar specimens 7 days old. figure 3. flexural and compressive strength of mortar specimens 28 days old. r e f a lg n a 1 a lg n a 5 a lg n a 10 d g 1 d g 5 d g 10 x g 1 x g 5 x g 10 c g 1 c g 5 c g 10 durability coefficient [–] 0.19 0.69 1.02 0.95 0.09 0.20 0.58 0.30 0.84 0.97 0.42 0.46 0.82 table 1. durability coefficient of tested mortars. 152 vol. 22/2019 physical-mechanical properties of hydraulic lime-based mortars figure 4. dg 1 and dg 5 (three of each from left to right) specimens after durability test. the specimens with the coefficient of 0.75 and higher are considered to be frost-resistant for the tested count of cycles, thus the algna 5 and 10, xg 5 and 10 and cg 10 specimens are frost resistant for 15 freezingthawing cycles. the values of durability coefficient of algna in the higher doses varied around 1, which indicates extreme durability in comparison with reference mortar, the coefficient higher than 1 can be explained as in the range of measurement errors. 4. conclusions the hardened physical-mechanical properties and durability of natural hydraulic lime mortars prepared from nhl 3.5 lime with addition of four different biopolymeric admixtures were studied with the following conclusions: • all the admixtures cause dosage-dependent flow value reduction; the most efficient ones are diutan gum and carrageenan while xanthan gum shows the highest dosage-dependency. • the addition of biopolymers increases the 7 and 28 days strength. diutan gum and carrageenan notably increase the 28 days flexural strength, while compressive one sustained similar to other two admixtures. • diutan gum addition worsens the durability of mortars, all other admixtures improve durability with dependency on growing dosage, the best performing is sodium salt of alginic acid, showing almost no strength reduction in higher doses. acknowledgements the work has been financially supported by the but specific research project fast-j-18-5274. references [1] e. knapen, d. van gemert. cement hydration and microstructure formation in the presence of water-soluble polymers. cement and concrete research 39:6–13, 2009. doi:10.1016/j.cemconres.2008.10.003. 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polytechnica ctu proceedings doi:10.14311/app.2019.24.0015 acta polytechnica ctu proceedings 24:15–20, 2019 © czech technical university in prague, 2019 available online at http://ojs.cvut.cz/ojs/index.php/app debris-fretting test of coated and uncoated zr-1%nb cladding marcin kopeća, ∗, martina maláa, ladislav cvrčekb, jakub krejčíc a research centre řež, hlavní 130, 250 68 husinec-řež, czech republic b czech technical university in prague, faculty of mechanical engineering, deparment of materials engineering, karlovo náměstí 13, 121 35 prague 2, czech republic c ujp praha a.s., nad kamínkou 1345, 156 10 prague – zbraslav, czech republic ∗ corresponding author: marcin.kopec@cvrez.cz abstract. debris fretting is one of the most significant fuel rod failure causes. an approach to mitigate this is to apply special coatings on the surface of the cladding that are supposed to increase the wear resistance. this paper summarizes the experiments performed in research centre řež on fuel cladding tube specimens. the tube segments made of uncoated and coated zr-1%nb alloy were exposed to fretting tests simulating the wire-debris vibrations impact on cladding during the reactor operation. the goal of these tests is to obtain results about the differences in wear resistance of uncoated and coated specimens. this paper shows the testing procedure in the air at room temperature and obtained impact markings on the specimens’ surfaces. keywords: cladding, fretting, nuclear fuel. 1. introduction the leaking fuel rods are an issue having been solved for a long time and remaining still relevant. in most cases the root cause of the failure is found which is visible in figure 1. nevertheless, in some causes the root cause cannot be identified at the power plant and then it is necessary to cool down the assembly in a spent fuel pool for few years to decrease its activity, then transport it to a hot cell facility, and perform detailed post-irradiation examination. this is associated with higher costs. the leaking fuel rod phenomenon is associated also with possible deteriorative handling and storage. this can stand as a challenge for utilities, who either have to invest in the repair of fuel assembly or loose space in their spent fuel pools. indirectly this issue affects also the fuel vendors when their fuel product is not so reliable and needs improvement. generally, the improvement can be beneficiary for the whole nuclear fuel market but the costs, to reach it, are excessive and may have been pointless. the nei report from 2016 specifies the leaking fuel rod costs as $1.5b per reactor [1] globally. one of the most important degradation mechanism of a nuclear fuel rod is the fretting wear coming from the flow-induced vibrations (fiv). the weak wear resistance of the cladding material can lead to its rupture. according to the nuclear news (in cooperation with epri), for the year 2010, the majority of failed fuel rod causes in the u.s. was related to the wear problem [2]. the more renowned mechanism of wear damage is grid-to-rod fretting. some studies and models have addressed this phenomenon [3–5], including power plant experiences [6]. on the opposite side there is the debris fretting failure mechanisms. the studies concerning the debris impact onto the fuel assembly are limited to new designs of anti-debris filters mounted in the fuel assembly bottom nozzles, e.g. filtering wires of diameter more than 1 mm or plates of the thickness of more 0.3 mm and lengths reaching more than 10 mm [7]. an attempt to limit the problem is the introduction of the foreign material exclusion (fme) policy, which is more and more incorporated by the nuclear power plants into their internal regulations. this, although, can proof ineffective in the case of object of diameter less than 1 mm and less than 10 mm in length, e.g. wires coming from the maintenance of the power plant. the fiv in those cases gets into high-frequency vibrations. hence the force load of a single debris-cladding contact is relatively very small, the numerous impacts can lead to fretting wear when the affected range is limited. another approach lowering the debris-impact onto the fuel cladding integrity is the introduction of coatings on the cladding surface, which are usually formed as a very thin film (microns) of highly wear-resistant material. this issue is directly related to the accident tolerant fuels (atf) concepts, being the matter of a research world widely. the most known is the crn layer coating which dates back to the 70’ and has its own history in e.g. halden reactor project (hrp) activities [8]. those days, the coatings and the process of their application, were extremely costly and ineffective. with the development of technology and the research constantly pending in this field, the coatings’ performance has been increased. similarly the range of potentially useable material for coatings. no ultimate material has been specified and in the aim of mechanical tests is to select the group of the 15 http://dx.doi.org/10.14311/app.2019.24.0015 http://ojs.cvut.cz/ojs/index.php/app m. kopeć, m. malá, l. cvrček, j. krejčí acta polytechnica ctu proceedings figure 1. the percentages of nuclear fuel rod failure causes in u.s. for 2010 [2]. most promising in the view of their endurance also under the load of debris in high-frequency vibrations. such process constitutes a qualification process for new atf materials and new coatings on standard cladding materials. the r&d tasks to specify the best accessible material is world widely ongoing the fuel vendors are equally involved into this process as the research organizations [9, 10]. research centre řež (cvř) has developed its own methodology for debris fretting of nuclear fuel cladding testing in dry conditions with the possibility to widen the spectrum for wet and high-temperature dry tests. in this way, it is possible to test and compare different types of atf claddings, including even sic/sic composites, which are supposed to withstand the debris impact even in elevated temperatures. in the cooperation with the czech technical university (ctu) in prague and ujp praha a.s. several types of coatings on standard zr-1%nb cladding were tested for fretting wear resistance. all tests included the reference on the as-received specimen of zr-1%nb cladding. this work presents the testing method and preliminary outputs of the materials’ comparison. 2. testing method of debris fretting wear as no standards have been found in the open literature for debris fretting wear testing as well as no models have been found for the prediction of the debris particle inside the reactor core during reactor operation, the authors decided to foreseen the potentially worst scenario. hence it is on the verge of possibilities to model the behavior of the small wire particle inside the reactor core, the case where one of the ends of the wire is caught inside a fuel assembly spacer grid was selected. in this way, the second end exerts a repetitive load on the cladding surface, depending on the fiv of the particular fuel assembly. as the anti-debris filters capture foreign objects of ca. 1 mm in each dimension [7], a stainless steel wire of the diameter of 0.4 mm was specified for the fretting tests. the testing equipment can house as many as 8 specimens including the reference one. it is based on the rotation cycling of the foreign object imitator (i.e. wire) with the possibly constant radial speed of ca. 1000 rpm. the precise number of contacts between the foreign object imitator and the specimens are recorded with voltage circle closed by the wire and the specimen. the wire has one free end that cyclically gets into smooth slide-contact with testing specimens. it is ensured by the bending of the used wire as can be seen in figure 3a. to avoid sliding of the wire over the surface of the specimen, it is kept between two clamps which are leaving no more space than twice the diameter of the used wire. in this way, the slide-spot of the wire over the specimen is at the same level as the fastening of the wire itself, what can be seen in figure 3b. the specimens are mounted symmetrically against the rotating shaft inside the testing equipment. the length of the wire is ca. 2 mm longer than the distance 16 vol. 24/2019 debris-fretting test of coated and uncoated zr-1%nb cladding figure 2. testing equipment with 2 specimens in parallel testing: right – uncoated zr-1%nb; left – crn-coated zr-1%nb. (a) schematic upper view – specimen (red), wire (blue), the arrow shows wire’s movement direction, black – pivot. (b) clamps (black) and wire positions. figure 3. the wire scheme. between the shaft and the specimen. in this way, the equal angle distances between the specimens and the length of the wire ensures almost equal force that the wire exerts on the specimens’ surfaces. what is more, the contact force is extremely low to keep the wire-specimen in the conditions that will enable free sliding of the wire against the specimen. the impact of additional coating layer thickness is due to the above negligible. to keep the reference between each batch of testing specimens and to preserve the correlation between the various types of coatings, 4-time ranges were specified: 120 min, 200 min, 350 min, and 500 min. the cycling of the specimens and the recording of the impact appearance is controlled from the computer. the equipment allows testing the specimen in wet condition up to 100 °c. while mounted in a special chamber, it is able to run a couple-of-days long experiment in temperatures reaching up to 800 °c (dry environment). the hot cells are able to house a long-lasting experiment (50 days) in the furnace in temperatures reaching 1100 °c (dry environment). in these hot-cell conditions, irradiated samples can be tested. the testing in elevated temperatures is important for the candidates of gen-iv cladding materials as sic/sic, which are meant to work in temperatures exceeding 800 °c. those materials should be able to withstand the conditions of possible accident, when the temperature could exceed 1000 °c [11]. 3. testing specimens of coatings for the testing, there were selected five specimens of different cr-based coatings, including the composite coatings. the coatings were prepared and applied at ctu in prague in cooperation with ujp praha a.s. the cr-based coatings were deposited on outer surfaces of zr-1%nb tubes by magnetron sputtering from cr (99.6 %) targets in the hauzer flexicoat 850 industrial system. pure metallic chromium coatings were prepared in the argon atmosphere and ceramic crn and crn2 coatings in a gas mixture of argon and nitrogen at the deposition temperature of about 250 °c. as the coatings to be tested, simple layers of pure chromium (cr), chromium nitride (crn) and super stochiometric chromium nitride crn2. as the mul17 m. kopeć, m. malá, l. cvrček, j. krejčí acta polytechnica ctu proceedings ticomponent coatings were testerd chromium nitride and chromium (crn+cr) and super stochiometric chromium nitride and chromium (crn2+cr). before deposition, the tubes were ultrasonically cleaned in acetone, ethanol, and demineralized water and dried with in furnace at 90 °c. the tubes were placed then hanged into the vacuum chamber on rotating holders in the flexicoat system where the tube surfaces were cleaned by ion etching in the argon plasma. this process removes the thin zr-oxide and other impurities thus improving adhesion of the coating. the deposition parameters vary depending on the particular coating. a thin metallic cr-layer was deposited on the substrate first when depositing ceramic coatings to improve adhesion. the thickness of the as-deposited coatings was measured with a calotest (csm, switzerland). the resulting thicknesses are presented from the outermost coating towards the substrate in table 1. coating layer 1 (µm) layer 2 (µm) total (µm) cr 18.6 – 18.6 crn 16.4 – 16.4 crn2 10.7 – 10.7 crn + cr 16.4 18.6 35.0 crn2 + cr 10.7 18.6 29.3 table 1. measured thicknesses of applied coatings. 4. the preliminary outputs at first, the tested specimens and their fretting marks were visually examined. they have not been thoroughly measured yet. they are prepared for replicamethod measurements with the use of high-resolution scanning – 5 micron precision. nevertheless, using the vertex automated precision measurement system it was possible to prepare good quality pictures of the fretting grooves. those preliminary outputs already can show the increase in wear resistance when the coatings basing on cr are applied. the tables below compare each batch of the tested specimens. however, the surface treatments of "as received" – uncoated and coated specimens are quite different, no additional improvement like polishing of coated specimens was applied. the tests were deliberately conducted in such conditions because of the fact, that any potential lta assembly will house either both uncoated and coated fuel rods or will have only coated rods but will be placed next to uncoated assembly. no information was found about the planned additional treatment and the supposition is that in the case of debris contact, the coated specimen can be polished "freely" with the debris. this is to verify the behavior of the mixed-core, which can ease the change from regular fuel to atf without the necessity to change the whole fuel batch at once. what is more, testing (a) sediments after testing of crn2+cr-coated zr1%nb. (b) sediments after testing crn-coated zr-1%nb. figure 4. sediments. the specimens in "as received" state – both coated and uncoated – gives more information about the inherent strength of the coating. in this way, the potentially ultimate candidates’ range can be limited. a very important output that can be distinguished in those preliminary outputs are the rest-overs coming from the friction between the wire and the specimen. their highest accumulation was observed by the crn2+cr composite coating and the crn coating. as no such accumulation of particles was visible in the vicinity of reference specimen and as the length of the wire did not alter after all the tests, it is supposed that the sediment origins in the coated specimen. the most probable is the coating wear in the contact area but without breaking the coating structure. figures 4a and 4b show sediment accumulation. the precise answer about the sediment can be delivered by microstructure analysis which is planned for the second half of 2019 – after the replica measurements. 5. conclusions the cvř has developed its own methodology and equipment for debris fretting testing of cladding materials. the testing equipment can be extended for tests in elevated temperature and long-lasting tests in temperatures exceeding 1000 °c. in the cooperation with ctu in prague and ujp praha a.s. five different types of cladding materials were tested in the same conditions for their comparison. the photos obtained with the vertex equipment clearly indicates the positive effect of coating application on the wear resistance of standard zr-1%nb cladding. on the reference uncoated "as-received" samples the grooves from 18 vol. 24/2019 debris-fretting test of coated and uncoated zr-1%nb cladding 120 min 200 min 350 min 500 min crn coated uncoated crn coated uncoated crn coated uncoated crn coated uncoated 120 min 200 min crn+cr coated cr coated uncoated crn+cr coated cr coated uncoated 350 min 500 min crn+cr coated cr coated uncoated crn+cr coated cr coated uncoated 120 min 200 min crn2+cr coated crn2 coated uncoated crn2+cr coated crn2 coated uncoated 350 min 500 min crn2+cr coated crn2 coated uncoated crn2+cr coated crn2 coated uncoated 19 m. kopeć, m. malá, l. cvrček, j. krejčí acta polytechnica ctu proceedings fretting increases with the number of loads, whereas, in the case of coated material, the zr-1%nb matrix seems untouched. high-precision measurements, using replica-method, are planned to specify the real depth of the obtained debris fretting marks. this activity will allow to compare the wear resistance of the coating materials and serve for the qualification reference of the coating in the case of its utilization in the nuclear fuel industry. the powder particle sediment was observed in the vicinity of coated specimens only, with its highest accumulation close to coatings made of crn2+cr and crn. the origin of the powder particles is supposed in the surface abrasion of the coating. to investigate the loss of material and friction abilities of the tested coating, the microstructural analysis in the area of fretting marks is planned for later this year. the next steps considered for these specimens are testing at elevated temperatures and testing in water environment. this will bring another data to complete the material resistance against fretting damage. the dry conditions testing is a first step to specify the most promising candidates for future atf cladding materials. more representative data for direct testing will be obtained from the wet test that can include also more aggressive and corrosive environment. acknowledgements financial support of this research through čez a.s. company is gratefully acknowledged. also, this work was supported by the ministry of education, youth and sport of the czech republic, programme npu1, project no lo1207. the presented work was financially supported by the ministry of education, youth and sport czech republic project lq1603 (research for susen). references [1] nuclear energy institute, nuclear cost in context retrieved march 2016 from: https://www.nei.org/ corporatesite/media/filefolder/policy/papers/ nuclear-costs-in-context.pdf?ext=.pdf, 2016. [2] the nuclear news interview, k. edsinger. epri and the zero fuel failure program fuel special edition, december 2010. [3] m. pettigrew. the vibration behavior of nuclear fuel under reactor conditions. nuclear engineering and design 114 pp. 179–189, 1992. [4] s. kovács, j. stabel, m. ren, b. ladouceur. comparative study on rod fretting behavior of different spacer spring geometries. wear 266 pp. 194–199, 2009. [5] j. yan, k. yuan, e. tatli, z. karoutas. a new method to predict grid-to-rod fretting in a pwr fuel assembly inlet region. nuclear engineering and design 241 pp. 2974–2982, 2011. [6] k.-t. kim. a study on the grid-to-rod fretting wearinduced fuel failure observed inthe16x16 kofa fuel. nuclear engineering and design 240 pp. 756–762, 2010. [7] j.-k. park, s.-k. lee, j.-h. kim. development of an evaluation method for nuclear fuel debris-filtering performance. nuclear engineering and technology 50 pp. 738–744, 2018. [8] j. h. jackson, d. paraventi, m. wright. proceedings of the 18th international conference on environmental degradation of materials in nuclear power systems – water reactors. portland, usa 2017. [9] l. hallstadius, s. johnson, e. lahoda. cladding for high performance fuel. progress in nuclear energy 57:71–76, 2012. [10] e. lahoda, f. franceschini. advanced fuels for future light water reactors. westinghouse electric company llc, pittsburgh 2011. [11] k. barrett, s. m. bragg-sitton, d. galicki. advanced lwr nuclear fuel cladding system development trade-off study. idaho national laboratory 2012. 20 https://www.nei.org/corporatesite/media/filefolder/policy/papers/nuclear-costs-in-context.pdf?ext=.pdf https://www.nei.org/corporatesite/media/filefolder/policy/papers/nuclear-costs-in-context.pdf?ext=.pdf https://www.nei.org/corporatesite/media/filefolder/policy/papers/nuclear-costs-in-context.pdf?ext=.pdf acta polytechnica ctu proceedings 24:1–6, 2019 1 introduction 2 testing method of debris fretting wear 3 testing specimens of coatings 4 the preliminary outputs 5 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0290 acta polytechnica ctu proceedings 33:290–294, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague expansion of sustainability in the aci 318 building code shana kelley kpff consulting engineers, 1601 5th avenue, ste 1600, seattle, wa 98101, usa correspondence: shana.kelley@kpff.com abstract. the american concrete institute (aci) has long recognized the growing importance of accounting for environmental sustainability in the design of concrete structures. as the next version of the aci 318 building code is being developed (aci 318-25), a new sustainability subcommittee, subcommittee n, has been formed and tasked with expanding the provisions related to sustainability in the code. this paper reviews the current provisions in the aci 318 code that are related to sustainability and explores the possible future expansion of code and commentary language to provide better guidance on sustainable design. areas of focus for expansion of code language, including definitions of sustainability metrics, as well as the defined baseline for comparison, are discussed. while sustainable design is increasingly being specified in concrete projects, the sustainable properties of concrete are often interrelated with other aspects of design, including resiliency and intended lifespan. approaches to acknowledging the impacts of life cycle considerations on the sustainability of concrete construction are also explored. finally, the anticipated timeline for the development of aci 318 sustainability provisions is described, as well as potential interactions with other code writing organizing that are currently developing sustainability provisions. keywords: aci 318, code, sustainability. 1. introduction historically, concrete building codes have focused mainly on provisions related to strength, serviceability, stability, and durability of the structure. however, as the priorities of policy makers, owners, and designers have increasingly included environmental sustainability, the question of the role of the concrete building code in setting standards related to sustainability has arisen. this paper includes a summary of how the current aci 318 building code addresses sustainability, an exploration of the current practices in the design and specification of sustainable concrete, and finally a description of the areas of expansion that are currently being worked on by the new aci 318-n sustainability subcommittee. 2. concrete sustainability at aci and in the aci building code as the focus on environmental sustainability of building materials has increased in recent years, the american concrete institute (aci) has expanded the committees, standards, and publications related to concrete sustainability. in particular, aci technical committee 130 sustainability of concrete, has been actively producing technical publications on approaches to the design, specification, and performance of sustainable concrete. there are also numerous committees studying concrete topics related to sustainable concrete including supplementary cementitious materials, alternative cements, and recycled aggregate. as the practice of incorporating sustainability into buildings has become more common, the need to incorporate provisions related to sustainability within the concrete building code has increased. provisions directly referencing sustainability were first incorporated in section 4.9 of the 2014 version of the american concrete institute building code, aci 318-14 [1]. these provisions, which remain essentially unchanged in the newly released aci 31819, acknowledge that design professionals are permitted to specify sustainability requirements, but do not define how sustainability of concrete could be specified or considered in the design of a concrete structure [2]. while the code language with direct references to sustainability remained unchanged in the aci 31819 code, there were several additions and expansions within that code edition that are associated with materials that are increasingly being used in projects targeting sustainability goals. these include new provisions regarding the use of alternative cements and recycled aggregates. the current code cycle for aci 318 is working toward further expanding the concrete sustainability provisions with the establishment of a new sustainability subcommittee, aci 318-n. this subcommittee is now working toward developing new provisions related to sustainability within the building code. 290 https://doi.org/10.14311/app.2022.33.0290 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 aci 318 building code sustainability expansion 3. how sustainability is being considered in design before addressing how the approach to defining sustainability can be incorporated into a concrete building code, it is important to understand how the current market is regulating environmental sustainability in concrete. 3.1. environmental sustainability in relation to cement content since the cement in concrete contributes the majority of the emissions in the production of concrete, the majority of environmental sustainable efforts have been associated with cement reduction. until recently, most design professionals and owners utilized either maximum amounts of cement or minimum amounts of cement replacement with supplementary cementitious materials to define sustainability. however, as information related to the environmental impacts of materials has become more readily available, the ways that the sustainability of concrete is being defined are changing. 3.2. a new focus on embodied carbon and environmental product declaration much of the historical focus on the sustainability of buildings has been on the reduction of the carbon utilized for the operations of buildings. thus, energy codes have increasingly required better insulated structures and more efficient mechanical systems. however, as these systems have improved and the operational carbon footprint of structures has reduced, the focus is now increasingly shifting to reductions of embodied carbon in buildings. embodied carbon is the global warming potential (gwp) associated with the manufacture of a product. in order to quantify the gwp of a building, design professionals are increasingly utilizing environmental product declarations (epds) to calculate the impacts of elements of a building. epds are reports, typically abiding by the requirements of the international organization for standardization (iso) 21930 for construction products [3], that quantify various environmental impacts of the production of a product. the declaration of environmental indicators includes environmental impacts, resource use, energy use, and waste flows. as shown in figure 1; gwp, ozone depletion, and freshwater use are amongst the impact categories that are typically included in the summary of impacts for a product. many industries have created industry-wide epds that reflect the average impacts of a product across an industry, but individual manufacturers are increasingly creating epds for their specific products. it should be noted that epds typically do not account for the full life cycle impacts of the use of a product, as their scope usually only includes the "cradle to gate" portion of the cycle, where the "gate" is figure 1. example of the summary of environmental impacts from an epd for a concrete mixture [4]. the gate of the manufacturing facility. so while epds can help quantify the impacts of the production of a product, the impacts from transportation to site, construction, use, deconstruction, and end-of-life are not included. this increased availability of information regarding the environmental impacts of concrete is starting to change the way that design professionals, building owners, and building officials are requiring and specifying the sustainability of concrete. 3.3. collection of epds into searchable databases as the number of epds has grown, numerous databases have been developed to aggregate the information in one place to better compare products. one of these online directories, the embodied carbon in construction calculator (ec3) has an extensive and growing collection of epds that it utilizes to calculate the gwp of building projects. as can be seen in figure 2, concrete represents by far the largest number of epds collected [5], likely due to the fact that each concrete mixture at a plant has a separate epd and because the concrete industry has rapidly adopted the use of epds. 291 shana kelley acta polytechnica ctu proceedings figure 2. graph showing number of epds for materials in ec3 carbon calculator software [5]. 3.4. life cycle assessment and resiliency in addition to the trend of examining the relative environmental impacts amongst a group of products within a certain industry, designers are increasingly examining the overall impacts of building construction and use over the entire life of the building. this process, called life cycle assessment (lca), is used to calculate the environmental impacts associated with all the life states of a product. often when lca is undertaken, the performance of a structure is compared to a baseline structure to calculate the degree to which the building has reduced its impacts. many sustainable certification systems for building construction require that an lca for a structure be performed, sometimes with a requirement for reductions of impacts or performance below a set threshold. while lca studies do assign a lifespan to certain products and adjust the overall impact of those products based on the selected lifespan of the structure, the resiliency of the structure is not necessarily accounted for with lca. with concrete structures, one of the benefits often cited is the resiliency of the structure in the event of natural disasters, extreme exposures, an extended lifespan, or other conditions. the question of how those qualities are to be accounted for is one that should be addressed when examining the life cycle impacts. 3.5. mandating of sustainability in local building codes with the increased focus on embodied carbon in building structures, local jurisdictions are beginning to mandate or consider mandating some characteristics of concrete related to sustainability. these ordinances and legislation vary in their scope and requirements, but many set limits on gwp and/or cement content. the advent of these laws lends an importance to ensuring that the concrete building code, aci 318, adequately addresses the items that must be considered when accounting for the sustainability of concrete structures. 4. aci 318-n sustainability subcommittee to address the increasing trend of defining limits on the environmental impacts of concrete in the design of buildings, the aci 318 building code has established a new sustainability subcommittee. this new subcommittee, aci 318 subcommittee n, is tasked with expanding the provisions related to sustainability in the concrete building code. the members on this subcommittee include practicing engineers, researchers, and concrete producers. many have extensive experience specifying and producing concrete to meet sustainability goals. 4.1. timeline for code development the aci 318-n subcommittee had its first official meeting at the october 2019 aci fall convention in cincinnati, ohio. the last published edition of the aci 318 building code was aci 318-19. it is anticipated that the next version released will be aci 318-25 and any provisions developed by the 318-n subcommittee that are approved by the main aci 318 committee will be included in the next edition. 292 vol. 33/2022 aci 318 building code sustainability expansion 4.2. possible areas of sustainability expansion in aci 318 while the manner and details surrounding the incorporation of sustainable concrete have not yet been established, the subcommittee is currently considering creating an appendix with sustainability provisions and commentary that will be part of the code. since the inception of the aci 318-n subcommittee, a number of task groups have been established to begin the process of writing code provisions related to sustainability. these task groups include the following: • scope section task group this group is addressing both the existing sustainability section of the code and its reference to the proposed sustainability appendix. • sustainable design of structural concrete task groups there are two separate task groups associated with this topic: ⊲ sustainable design this task group is focused on how the design of structures should account for concrete when sustainability is an objective. in particular this group will be looking at approaches such as life cycle assessment and how aspects of concrete should be accounted for. ⊲ resiliency this task group is examining how resiliency of concrete structures relates to the consideration of the sustainability of a structure and how it should be considered. this group is also examining resiliency in general and how it is accounted for in the code. • concrete mixture design information task group this group is working to develop provisions that outline how concrete mixtures can be specified to require sustainability metrics. the group is currently examining several approaches and is debating what methods to include. the interface between this section and the sustainable design section is still under consideration. • recycled aggregate task group this group is looking at the new provisions in aci 318-19 related to the use of crushed hydraulic-cement concrete and recycled aggregate. in particular, the new provisions require testing and quality control programs for these types of aggregate [2] and the task group is examining whether additional guidance could be provided in the commentary of this section regarding those programs. this code section falls within the overview of a separate aci 318 subcommittee, thus any changes to this section will require coordination across subcommittees. • cements and alternative cements task group chapter 26 of aci 318-19 contains a list of acceptable astm specified cementitious materials along with a new section regarding alternative cements [2]. this task group is examining whether additional cementitious materials with astm specifications should be added and whether any additional clarifications are needed for the alternative cement section. as with the recycled aggregate topic above, any changes to the provisions related to cement will need to be coordinated across aci 318 subcommittees. since the aci 318 code has both code language and commentary, these groups will be writing proposed sections for both. the commentary will likely include discussions of what approaches could be taken to achieve the code standards. as this subcommittee’s work progresses, the relationships between the work of the different task groups will emerge and there will be more clarity regarding the form and content of the proposed provisions. 4.3. interactions with other sustainable concrete code writing organizations as the aci 318-n subcommittee is beginning its work to expand the code language related to sustainability in the aci 318 building code, the subcommittee is reviewing the work of other concrete code writing bodies and their approach to the incorporation of sustainability into their codes. in particular, aci 318-n has studied the sustainability sections of the fib model code 2010 and sought insight into the active developments in fib commission 7: sustainability (com7). in order to further foster coordinating and information sharing, aci 318-n will have a joint meeting with com7 at the fib international conference on concrete sustainability that will occur in prague in september 2020. 5. conclusion this paper has addressed how the current aci 318 building code addresses sustainability, the changing landscape of sustainable design and specification, and the possible areas of expansion within aci 318. the efforts related to this expansion are in their infancy, but with the wide use of aci 318 the possible impacts are great. the implementation of sustainable regulations on concrete are growing and will continue to grow regardless of the incorporation of sustainability in this building code. however, if efforts are made to make new provisions pertinent and easily implementable, the potential to make specifying sustainable concrete more consistent and logical are significant. acknowledgements the author gratefully acknowledges the travel support for this conference provided by the american concrete institute technical cooperative representative program. the opinions expressed in this paper are the authors own and do not necessarily represent the opinions of aci. 293 shana kelley acta polytechnica ctu proceedings references [1] american concrete institute. building code requirements for structural concrete (aci 318-14) commentary on building code requirements for structural concrete (aci 318r-14) ed aci committee 318, farmington hills, mi: american concrete institute, 2014. http: //aghababaie.usc.ac.ir/files/1506505203365.pdf. [2] american concrete institute. building code requirements for structural concrete (aci 318-19) commentary on building code requirements for structural concrete (aci 318r-19) ed aci committee 318, farmington hills, mi: american concrete institute, 2019. https://www.usb.ac.ir/filestaff/ 5526_2020-1-25-11-12-7.pdf. [3] iso 21930:2107. sustainability in buildings and civil engineering works core rules for environmental product declarations of construction products and services, geneva, international organization for standardization, 2017. [4] astm international. central concrete environmental product declaration mix 3daz75z2 south san francisco (wet) plant, 2020. https://www.astm.org/c ertification/docs/epd_central_concrete.html. [5] building transparency. embodied carbon in construction calculator, 2020. https://www.buildingtransparency.org. 294 http://aghababaie.usc.ac.ir/files/1506505203365.pdf https://www.usb.ac.ir/filestaff/5526_2020-1-25-11-12-7.pdf https://www.astm.org/certification/docs/epd_central_concrete.html https://www.buildingtransparency.org doi:10.14311/app.2021.30.0053 acta polytechnica ctu proceedings 30:53–57, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app fracture energy of illitic rammed earth with high water-clay ratio barbora mužíková∗, tereza plaček otcovská, pavel padevět czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: barbora.muzikova@fsv.cvut.cz abstract. the article is focused on design of mixture of rammed earth, producing, testing and determination of fracture energy of unfired rammed earth and its stress-strain curve in tensile bending test. three different mixtures of rammed earth were designed and tested. the amount of water and binder is one of the key properties of the rammed earth, the amount of the water is expressed by the water-clay ratio. mechanical properties of the earth material highly depend on the composition of sand, clay and water. the prescription agl iii with 80 % of sand, 20 % of clay and 0.400 water-clay ratio reached the maximum value of fracture energy 4.858 ± 0.002 j/m2 and set agl v had the minimum value 1.934 ± 0.310 j/m2. keywords: fracture energy, illite, rammed earth, stress-strain curve, water-clay ratio. 1. introduction the rammed earth is nowadays considered as a sustainable way for modern building. the article is focused on illitic rammed earth with high water-clay ratio. the mechanical property that was searched is the fracture energy and stress-strain curve. earth as a building material is coming back thanks to the trend to use this natural and traditional material of our ancestors. an expression “sustainable building” describes this trend and unfired earth perfectly meets all its principles. besides that it is also an environmentally friendly final product which is not synthetically fabricated. moreover, it is recyclable and there is no need after demolition to take this material to the garbage dump. but the main disadvantage for its using is that the mechanical properties needed for a structure design are not well known. this problem is more complicated because the mechanical properties depend on a composition of the used earth [1–5]. 2. composition of rammed earth as written before, the mechanical properties depend on the composition of earth. earth is a mixture of clay, sand and water. from the point of mechanics, the clay is a binder (such as cement for concrete), sand is a filler that makes the structure and water activates the binding forces of clay. 2.1. binder the clay is a binder for rammed earth. unfortunately for searching the mechanical properties, there are three basic types of clay – kaolinite, illite and montmorillonite. they differ in chemical composition. clay minerals are hydrous aluminium phyllosilicates with variable amounts of iron, magnesium, alkali metals, alkaline earths. the structure is made by sheets. the number and size of these sheets and gaps between them have an impact on water absorption of the final product. the clay changes the properties of the final product – mechanical properties and the colour is different as can be seen in the fig. 3. the specimens tested for the fracture ennergy determination were made with illite clay. 2.2. filler the mixture of clay and sand (or sand with bigger parts like gravel) is called a loam. the process of the formation is called an erosion of rocks. it takes a very long time and this process occurs naturally and depends on local weather conditions and surrounding nature. as a result of erosion, the particles of different sizes arise – clay(less than 0.002 mm), silt (0.002 – 0.060 mm), sand(0.060 – 2.000 mm) and gravel (more than 2.000 mm). the smaller parts have the function of a binder and the others have the function of a filler. 2.3. water-clay ratio water is the starter for binding reaction of clay. thus it is very important to have the suitable amount of water. the amount of water in rammed earth is expressed by the water-clay ratio. the meaning is the same as the water-cement ratio at concrete [6–12]. 3. production of specimens the principle of building rammed earth using framework is shown in fig. 1 and the production in laboratory in fig. 2. first of all, the framework is built to create the shape of the future wall. secondly, a small part of the earth mixture is put in the framework 53 http://dx.doi.org/10.14311/app.2021.30.0053 http://ojs.cvut.cz/ojs/index.php/app b. mužíková, t. plaček otcovská, p. padevět acta polytechnica ctu proceedings figure 1. procces of building earth wall. figure 2. prefabrication of the specimens in laboratory conditions. figure 3. specimens with different type of clay as a binder. 54 vol. 30/2021 fracture energy of illitic rammed earth set sand/clay ratio water-clay ratio pcs of specimens bulk density type of clay [%/%] [−] [−] [kg/m3] [−] agl iii 80/20 0.400 3 2103 illite agl v 85/15 0.370 3 2005 illite agl ix 80/20 0.450 3 2107 illite table 1. composition of designed mixture for rammed earth. and compressed. then next layers of rammed earth are put and compressed and together they create the rammed earth wall. as the final step, the framework is removed. in the same way the specimens are made in laboratory, but there is no big framework for the wall but just small ones to create the specimen of exact dimension. the difference is the step before the producing. the prescription has to be chosen. that means to choose a type of clay, then a ratio between clay and filler and a water-clay ratio. the accurate amount of clay, the sand and water are weighed and mixed. for the need of three-point bending test three different prescriptions were designed and tested. the agl sets iii, v and ix are all with illitic clay. the composition of these set is in-depth described in the table 1. for the three-point bending test were made specimens of size 40×40×160 mm. another comperative tests were made in [11–13]. the photograph captured during the test is shown in the fig. 4. figure 4. the three-point bending test on specimen from rammed earth material. 4. method of determining fracture energy to get fracture energy the three-point bending test needs to be done. the force that is apllied on the specimen and the matching displacement are recorded. the force-displacement curve (f-δ) is recorded and the stress-strain curve (σ-#)can be calculated from these data. the get stress-strain we have to caltulate the strain: figure 5. the stress-strain curve for agl iii specimen numb. 3. # = δ ÷ l, (1) where l is the length of the specimen between the support points in the test. then the stress can be calculated: σ = f ÷ a, (2) where a stands for the cross area. now we have data to get stress-strain curve (fig. 5). next step is to calculated the fracture energy. the fracture energy is defined as the amount of energy necessary to create one unit area of a crack. firstly, the work of the apllied force wf is calculated (7): wf = ! δmax 0 f dδ , (3) and then the fracture energy is: gf = af ÷ bw, (4) where bw stands for the area of the crack (w is the length of the ligament and b is the width of the specimen). in this case this area is the same as the cross area a of the specimen. the schema of the test can be seen in the fig. 6. 55 b. mužíková, t. plaček otcovská, p. padevět acta polytechnica ctu proceedings figure 6. schema of the three-point bending test. set s/c w type of clay gf ft e [−] [−] [j/m2] [mpa] [mpa] agl iii 80/20 0.400 illite 4.858±0.002 0.059±0.002 68.742±28.224 agl v 75/25 0.370 illite 1.934±0.310 0.028±0.003 49.056±15.171 agl ix 75/25 0.450 illite 4.493±0.521 0.051±0.001 27.038±7.597 table 2. evaluated fracture energy, tensile strength in bending and modulus of elasticity. figure 7. the work to create the crack is the area under f -δ curve. 5. results of the measurements the result of the measurement were the arithmetic mean and the standard deviation for each set from the measured values, they are in table 2. there are the compositions of the mixture, used type of clay, evaluated fracture energy (± standard deviation), their tensile bending stregths (± standard deviation), elasticity modulus (± standard deviation). the values of fracture energy are shown also in fig. 8. set agl iii with water-clay ratio 0.400 has the highest value of the fracture energy 4.858±0.002 j/m2, agl ix with watter-clay ratio 0.450 just a little bit lower value 4.493 ± 0.521 j/m2 and set agl v has significantly figure 8. fracture energy of tested mixtures. lowest value 1.934 ± 0.310 j/m2. it is interesting that set agl iii with 20 % clay and water-clay ratio 0.400 have very simillar result as agl ix with 25 % clay and water-clay ratio 0.450. so the set agl ix have more clay and also more mixing water. also their tensile strengths are very similar 0.059 ± 0.002 and 0.051 ± 0.001. 6. conclusions three different mixtures were designed, producted and tested. measured data were evalueted a compared in relation to its composition. the paper was focused on evaluating fracture energy and strainstress curve of different sets of rammed earth. the mechanical properties of the earth material highly depend on the composition of sand, clay and water. 56 vol. 30/2021 fracture energy of illitic rammed earth the amount of water and the binder is one of the key properties of the rammed earth, it is similar to concrete, so the amount of water is expressed by the water-clay ratio. the water-clay ratio 0.400 seems to be optimal with the mixture agl iii (80 % sand and 20 % clay) and also very good results have the composition with water-clay ratio 0.450 for mixture agl ix (75 % sand and 25 % clay). list of symbols e elascity modulus [mpa] f load force [kn] gf fracture energy [j/m2] fc compressive strength [kpa] ft tensile bending strength [kpa] δ displacement [mm] ε strain [–] # bulk density [kg m−3] σ stress [kpa] l length between supports in the test [mm] a cross area [mm2] acknowledgements this research was financial supported by the czech science foundation (gacr no.18-0884s) and by the faculty of civil engineering at czech technical university in prague (sgs project no.16/201ohk1/3t/11). references [1] b. little, m. morton. building with earth in scotland: innovative design and sustainability, vol. 1 of first printing. scottish executive central research unit, edinburgh, 2001. [2] j. norton. building with earth, a handbook, vol. 1. intermediate technology development group limited, london, 1986. [3] g. minke. building with earth design and technology of sustainable architecture, vol. 2. de gruyter, berlin, 2009. [4] i. zabickova. hlinené stavby. era 21, brno, 2002. [5] m. rauch. refined earth construction and design with rammed earth, vol. 1. detail, munich, 2015. [6] kianfar, ehsan, v. toufigh. reliability analysis of rammed earth structures. construction and building materials (127):884–895, 2016. [7] h. araki, j. koseki, t. sato. tensile strength of compacted rammed earth materials. soils and foundation (56(2)):189–204, 2016. [8] l. miccoli, d. v. oliveira, r. a. silva, u. muller. static behaviour of rammed earth: experimental testing and finite element modelling. materials and structures (48(10)):43–56, 2015. [9] q. liu, t. l. tong. engineering properties of unstabilized rammed earth with different clay contents. journal of wuhan university of technology: materials science (32(4)):914–920, 2017. [10] r. el-nabouch, q. bui, o. ple, p. perrotin. characterizing the shear parameters of rammed earth material by using a full-scale direct shear box. construction and building materials (171):414–420, 2018. [11] q. bui, j. morel, b. v. reddy, w. ghayad. durability of rammed earth walls exposed for 20 years to natural weathering. building and environment (44(5)):912–119, 2009. [12] x. dong, v. soebarto, m. griffith. strategies for reducing heating and cooling loads of uninsulated rammed earth wall houses. energy and buildings (77):323–331, 2014. [13] t. otcovska, p. padevet. dependence of tensile bending strength of rammed earth on used clay composition and amount of mixture water. modern methods of experimental and computational investigations in area of construction ii pp. 48–53, 2017. 57 acta polytechnica doi:10.14311/app.2020.27.0079 acta polytechnica 27(0):79–83, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app relation between hardness of (ti, al)n based multilayered coatings and periods of their stacking mehdi zaouia, b, alexandre bourcereta, yves gaillarda, sylvain giljeanc, ∗, christophe rousselota, marie-josé pacc, fabrice richarda a université bourgogne franche-comté, institut femto-st, cnrs/ufc/ensmm/utbm, 32 avenue de l’observatoire, 25000 besançon, france b institut de recherche technologique, matériaux, métallurgie, procédés, 4 rue augustin fresnel, 57070 metz, france c université de haute alsace, lpmt, 2 rue des frères lumière, 68093 mulhouse, france ∗ corresponding author: sylvain.giljean@uha.fr abstract. this study aims to model, by using a finite element method, the relationship between the hardness and the period λ of metal/nitride multilayer coatings (ti0.54al0.46/ti0.54al0.46n)n in order to understand the increase of the hardness at the low periods [1] and then optimise the multilayer coating architecture to obtain the best mechanical properties. a 2d axisymmetric finite element model of the berkovich nanoindentation test was developed. the coating was designed as a stacking of ti0.54al0.46 and ti0.54al0.46n nanolayers with, in the first hypothesis, equal thickness and perfect interface. the elastoplastic behaviours of the metal and nitride layers were identified by berkovich nanoindentation experiments and inverse analysis on thick monolayer samples. the indentation curves (p-h) obtained by this model depend on the period λ of the stacking. simulated (p-h) curves were compared with experimental data on 2 µm thick films with different periods λ ranging from 10 to 50 nm deposited by rf magnetron sputtering using reactive gas pulsing process (rgpp). the model forecasts are very consistent with the experience for the largest period but the model does not reproduce the hardness increase at the lowest periods. the λ = 10 nm coating was analysed by electron energy loss spectroscopy (eels) on a transmission electron microscope. results show intermixing of the layers with the presence of nitrogen atoms in the metal layer over a few nanometers [1]. it was concluded that the metal/ceramic interface plays an important role at low periods. the addition in the model of a transition layer in the metal/nitride stacking, with an elastoplastic metal/ceramic medium behaviour, allows to reproduce the nanoindentation experimental curves. the thickness of this transition layer deduced from model updating method is in very good agreement with eels observations. keywords: finite element method, hardness, indentation modulus, nanoindentation, tial/tialn multilayer films. 1. introduction titanium and aluminium nitrides (tial)n coatings are extensively used in the industry since many years for their hardness, toughness and wear and oxidation resistances [2, 3]. it has been shown that the indentation modulus and the hardness of ti1-xalxn coatings depend on the aluminium content and that the optimal properties are obtained when the maximum al content is substituted for ti in face centred cubic structure close to x = 0.46 [1, 4]. one way to improve the hardness of that kind of coatings consists in depositing multilayer films. in this study, multilayer coatings are studied and consist in metal/ceramic stacking with alternate deposition of ti0.54al0.46 and ti0.54al0.46n nanolayers with three different periods of the bilayer metal/ceramic λ = 50, 16, 10 nm. in order to facilitate the reading of the paper, ti0.54al0.46 and ti0.54al0.46n are respectively written as tial and tialn. the main objective of the paper is to study the relationship between hardness of as-design multilayer films and periods of their stacking to understand the elastoplastic phenomena that are involved in the hardness increase experimentally observed at the lowest periods [1]. a 2d axisymmetric model of the berkovich nanoindentation is developed. then a method to identify the elastoplastic mechanical properties of the monolayer tial and tialn is proposed. the results are introduced in a multilayer stacking model that allows to obtain the indentation curves (p-h) as a function of the period λ. as a first hypothesis the interface between tial and tialn layers is considered as perfect without thickness. then a transition layer is added and allows to reproduce the experimental indentation curves. 79 http://dx.doi.org/10.14311/app.2020.27.0079 http://ojs.cvut.cz/ojs/index.php/app m. zaoui, a. bourceret, y. gaillard et al. acta polytechnica 2. materials and methods tial and tialn monolayer thin films were deposited onto si (100) wafer by rf magnetron sputtering using one sintered titanium/aluminium 66/33 at.% alloy target as described elsewhere [4]. same experimental conditions were used to deposit (tial/tialn)n multilayer films with different periods using reactive gas pulsing process (rgpp) [5] with a controlled pulsing flow rate of the nitrogen reactive gas as described elsewhere [1]. finally, the samples considered in this paper are summarised in table 1. sample period thickness λ (nm) ep (µm) tial / 3.0 tialn / 3.0 50 2.0 (tial/tialn)n 16 2.0 10 1.8 table 1. sample characteristics. nano-indentation tests were performed using an ultra nano-indentation hardness tester provided by anton paar. experiments were conducted using a berkovich tip and the multicycle procedure in load control up to 8 mn [1, 4]. a 2d axisymmetric finite element model of the berkovich nanoindentation test was developed with the ansys finite element software as described on figure 1. an implicit method with non-linear geometric formulation has been used with higher order 2d elements. a deformable indenter, with isotropic linear elastic law, was modelled with a half contact angle of 70.3° and a tip radius to fit the experimental tip defects. a friction coefficient of 0.1 was imposed at the contact between the indenter and the film. the mesh density has been refined so that the smallest elements are localised in the contact region. the film can be either monolayer tial or tialn or multilayer (tial/tialn)n. for the monolayer tial and tialn, isotropic elastoplastic behaviour laws were considered with linear hardening. elastic properties were obtained from experimental nanoindentation and plastic properties were estimated using finite element model updating. the approach used to extract the constitutive parameters of plastic law of tial and tialn consists in an optimisation problem; the objective of which is to minimise a cost function ω by using an algorithm based on the levenberg-marquardt [6, 7] method using the mic2m software [8]: θ̂ = argmin θ ω (p (t; θ) ,pexp (t)) (1) where θ̂ is the vector of the unknown parameters and ω the cost function i.e., the gap between pexp (t) the figure 1. description of the 2d axisymmetric finite element indentation model. experimental data and p (t; θ) the numerical results. then the cost function can be defined as: ω(θ) = 1 2t t∑ k=1 ( pk(θ) −pexpk p exp max )2 (2) where t is the number of acquisition times of the load pk(θ) = p (tk; θ) and pexpmax the maximum value of the experimental indentation load. 3. results and discussion 3.1. mechanical properties of tial and tialn monolayers all the experimental data of the indentation tests performed on the tial and tialn monolayer films have already been published elsewhere [1, 4]. hardness and indentation modulus values of the metal tial are significantly lower than the ones of nitride tialn. by postulating the poisson’s ratio as known, the elastic properties of the monolayer films can be deduced from the experimental data. then, plastic parameters were numerically deduced from the optimisation approach described in method section. results are presented in table 2. the comparison of the experimental and model indentation load curves are plotted in figure 2 and show a good agreement. 3.2. simulation of (tial/tialn)n multilayer films with perfect interfaces all the experimental data of the indentation tests performed on the multilayer (tial/tialn)n films have already been published elsewhere [1]. interesting observation is that both the hardness and the elastic modulus of the multilayer films follow a hall and petch law as the values linearly increase as a function of 1/ √ λ. in a first approach, the multilayer film was modelled as a stacking of nanolayers of tial and tialn with a perfect interface without thickness as showed in figure 3.a. the model was set for each 80 vol. 27/2020 young modulus poisson ratio yield stress linear hardening (gpa) (gpa) (gpa) tial 199 0.25 0.88 56.6 tialn 417 0.21 5.70 405.0 table 2. elastoplastic parameters of the metal tial and the nitride tialn. figure 2. loading indentation curves (p-h) of (a) tial and (b) tialn monolayers. figure 3. modelling of multilayer films, (a) with perfect interfaces (b) with a transition layer. figure 4. comparison of the (p-h) experimental and simulated curves for λ = 10, 16 and 50 nm with a perfect interface between tial and tialn nanolayers. stacking period λ = 10, 16 and 50 nm. a convergence study was carried out with 1, 2 or 3 elements in the thickness of the nanolayers. since the relative variation of the load is less than 0.5%, only one element, with a λ/2 size, was finally used in the thickness (figure 3.a), in order to minimise the calculation time. results of the as simulated (p-h) curves are plotted on figure 4. for the largest period λ = 50 nm, the model is in good agreement with the experience. furthermore, the gap between the experimental curves and the simulated ones increases when the period decreases to become important for the smallest period. therefore, the phenomenon that generates an increase in hardness when period decreases is not taken into 81 m. zaoui, a. bourceret, y. gaillard et al. acta polytechnica figure 5. for λ = 10 nm, (a) comparison of the (p-h) experimental and simulated curves with a transition layer, (b) eels results. account in the coating model without interfaces. 3.3. simulation of (tial/tialn)n multilayer films with a transition layer in the stacking, between each nanolayer of tial and tialn, a transition layer was added (figure 3.b.). for this transition layer, an elastoplastic metal/ceramic medium behaviour was considered. each elastoplastic parameter of the transition layer was assumed to be an average (mixture law 50/50) of the parameters of the tial and tialn nanolayers. the thickness of this transition layer was optimised by minimising the gap between experimental and simulated data as previously described. result of the simulated (p-h) curve with a transition layer is plotted on figure 5.a and show a good agreement between experimental and simulated curves for the lowest period. for λ = 10 nm, the simulated thickness of the transition layer, eptr = 2.12 nm, is close to the one of tial and tialn layers. this result is in very good agreement with the thickness of interfaces observed between tial and tialn layers evaluated by electron energy-loss spectroscopy (figure 5.b). 4. conclusion the elastoplastic properties of ti0.54al0.46 metal and ti0.54al0.46n ceramic monolayers were identified by comparison of berkovich nanoindentation experimental data and 2d axisymmetric finite element simulations. the mechanical properties of the monolayers were used in a multilayer model to forecast the mechanical properties of (ti0.54al0.46/ti0.54al0.46n)n multilayer coatings, with various periods, designed by rf magnetron sputtering using rgpp technique. the multilayer model was considered, in first hypothesis, as a stacking of tialn and tial nanolayers with equal thickness and perfect interface. that model allows to reproduce the indentation curves of the highest 50 nm period films. to fit the experimental indentation curve at the lowest periods, a transition layer was introduced in the model between each metal/ceramic interface. using an optimisation process of the model, this transition layer was found to be 2.12 nm thick for the lowest period λ = 10 nm, which is in good agreement with the results of electron energy-loss spectroscopy experiments. acknowledgements this work was performed in the frame of the research project resem 2019, managed by the institut de recherche technologique matériaux métallurgie procédés (irt m2p) and financially supported by the french program plan d’investissement d’avenir (pia). references [1] m.-j. pac, y. pinot, s. giljean, et al. investigation of ti0.54al0.46/ti0.54al0.46n multilayer films deposited by reactive gas pulsing process by nano-indentation and electron energy-loss spectroscopy. thin solid films 634:96 – 106, 2017. doi:10.1016/j.tsf.2017.05.015. [2] w. münz. titanium aluminium nitride films: a new alternative to tin coatings. journal of vacuum science & technology a 4:2717–2725, 1986. doi:10.1116/1.573713. [3] s. paldey, s. deevi. single layer and multilayer wear resistant coatings of (ti,al)n: a review. materials science and engineering: a 342(1):58 – 79, 2003. doi:10.1016/s0921-5093(02)00259-9. [4] m.-j. pac, s. giljean, c. rousselot, et al. microstructural and elasto-plastic material parameters identification by inverse finite elements method of ti(1-x)alxn (0<x<1) sputtered thin films from berkovich nano-indentation experiments. thin solid films 569:81 – 92, 2014. doi:10.1016/j.tsf.2014.07.037. [5] n. martin, j. lintymer, f. sthal, et al. procédé de pulvérisation réactive à signal de commande cyclique et dispositif correspondant 2008. [6] k. levenberg. a method for the solution of certain non-linear problems in least squares. quarterly of applied mathematics 2(2):164–168, 1944. 82 https://doi.org/10.1016/j.tsf.2017.05.015 https://doi.org/10.1116/1.573713 https://doi.org/10.1016/s0921-5093(02)00259-9 https://doi.org/10.1016/j.tsf.2014.07.037 vol. 27/2020 [7] d. w. marquardt. an algorithm for least-squares estimation of nonlinear parameters. journal of the society for industrial and applied mathematics 11(2):431–441, 1963. doi:10.1137/0111030. [8] f. richard. mic2m software: modeling and identification of the mechanical behavior of materials, 2000. 83 https://doi.org/10.1137/0111030 acta polytechnica 27(0):79–83, 2020 1 introduction 2 materials and methods 3 results and discussion 3.1 mechanical properties of tial and tialn monolayers 3.2 simulation of (tial/tialn)n multilayer films with perfect interfaces 3.3 simulation of (tial/tialn)n multilayer films with a transition layer 4 conclusion acknowledgements references doi:10.14311/app.2021.30.0081 acta polytechnica ctu proceedings 30:81–86, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app numerical comparison of transgingival and subgingival dental implants in regard to their stress distributions luboš řehouneka, ∗, aleš jíraa, gabriela javorskáa, daniel bodlákb a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b prospon spol. s.r.o., jiřího voskovce 3206, 272 01 kladno, czech republic ∗ corresponding author: lubos.rehounek@fsv.cvut.cz abstract. most modern dental implants differentiate in regard to the fixation of the abutment into two main categories the external or internal hexagon or octagon. we performed mechanical tests according to the čsn en iso 14801 standard on a dental implant variant using the external hex. we found that failure of all implant specimens occured below the screw head. to improve the current geometry, we performed numerical analysis of an alternative variant (internal hex) and compared it with analysis of the current geometry (external hex). it was found that the stress distribution of the variant with internal hex is preferable to the old variant. although extreme values of shear stress in the corresponding plane of loading are higher, they do not concentrate below the screw head, where the screw itself is thinner and more prone to breaking. therefore, it seems that the new variant of the dental implant is stronger, which is still to be proven by mechanical tests. keywords: external hexagon, fem, implant-abutment interface, internal hexagon, stress. 1. introduction the implant-abutment connection is one of the most important interfaces of a dental implant. it provides a means to apply torque required to screw the implant into bone and also serves as a connection between the intraosseous part and the abutment. there are many requirements and functions of this interface that may be less or more propagated in regard to the form of connection (internal or external). however, one of the most important functions is the prevention of rotation between individual parts [1]. this particular property is greatly influenced by the character of the interface. from a mechanical standpoint, there are several geometrical solutions among dental implants in regard to their placement and the fixation of the abutment. they include the internal morse taper, external morse taper, internal lobing, tube-to-tube systems and external and internal polygons. these last 2 main categories comprise the vast majority of clinically used implants as well as the majority of market share. the most common variants of the polygonal fixation are the hexagon or the octagon. our study includes 2 different variants of the polygonal variant – the external hexagon (transgingival variant) and the internal hexagon (subgingival variant). their names suggest their placement in the bone, as shown in fig. 1. historically, the first manufactured titanium dental implants by brånemark had the external hexagonal connection. it is a variant that is very simple a robust, however, it has some deficiencies, such as insufficient anti-rotational resistance or insufficient resisfigure 1. comparison of placement of the 2 analysed dental implant variants inside the bone. left – the internal hexagon variant (subgingival placement) and right – the external hexagon variant (transgingival placemenent). the lenght l indicates the portion of the intraosseous part that exceeds the bone after implantation. tance to intraoral forces with greater offset from the connection (e.g. top of the crown). to improve lateral and rotational stability, many other variants were developed (internal hex, spline connection, morse taper, tapered hexagon and other variants) [2]. 1.1. motivation for development the literature suggests that internal connections have mechanical advantages. it was found that the bending moments of zirconia implants are superior when paired with internal connection, compared to exter81 http://dx.doi.org/10.14311/app.2021.30.0081 http://ojs.cvut.cz/ojs/index.php/app l. řehounek, a. jíra, g. javorská, d. bodlák acta polytechnica ctu proceedings figure 2. documentation of the old external (eh) connection variant of the implant.the intraosseous part is the subject of a patent [7]. nally connected abutments [3, 4]. research suggests that external connections lead to greater stresses inside the components of the implant, mainly at the connection interface [5]. also, they pose as an alternative to external connections, which have proven to break in the abutment screw [6]. the abutment screw can be identified as a very weak component of the system as it is thin while being exposed to shear forces during mastication. the motivation to develop other variants of connection also stems from the fact that external connections have proven to decrease levels of crestal bone at the peri-implant area after 1 year of loading and also exhibit greater bone remodelling and bone loss when compared to internal connections within the same set of circumstances [8]. furthermore, internal abutment connections have better lateral stability as the lever arm of forces is shorter, thus reducing the value of maximum force that can be applied in the lateral direction. according to literature, this results in better stress distributions, less micromovement and higher survival probability [9]. other factors, such as surface treatment, anodization, implant length, abutment material etc. also influence bone loss. another aspect of potential improvement over the external connection types is improving the transfer of loads in the surrounding bone and minimizing the gaps at the implant-abutment interface which may figure 3. documentation of the new internal (ih) connection variant of the implant.the intraosseous part is the subject of a patent [7]. lead to bacterial colonization [10] and subsequent biofilm accumulation [11]. 2. materials and methods the means of comparison in this study are fem analyses of 2 selected dental implant variants in regard to their stress distribution. the selected variants are the external hexagon (eh) and the internal hexagon (ih), as shown in fig. 2 and fig. 3. the simulations were performed on both variants on the basis of conditions and results of mechanical tests of the eh variant so as to verify the viability of the newer ih variant in regard to stress distribution. 2.1. mechanical tests we performed mechanical tests on the old eh variant according to the iso 14801 dentistry–implants– dynamic fatigue tests for endosseous dental implants standard [12] on a total of 6 specimens of dental implants (fig. 4). the tests were performed on a double-column pneumatic press litem vertical double column vdc a-6. the machine complies with iso 7500-1 [13] and enables for constant crosshead displacement. a sample p-d diagram of the tested eh specimens is shown in fig. 6. all specimens were embedded in the dentacryl fixating medium (methylacrylate resin for technical use). the resin has good mechanical properties and thermal insulation properties. its bending strength is 50 mpa, tensile strength is 20 82 vol. 30/2021 comparison of transgingival and subgingival implants figure 4. a disassembled specimen of the eh variant used for the mechanical tests. although the body of the implant is different, the locking system including the abutment and screw are the same. specimen number force [n] 1 1278.7 2 1387.6 3 1301.2 4 1486.0 5 1275.8 6 1233.2 mean value 1327.1 table 1. maximum static load force for all eh specimens. mpa and compressive strength is 80 mpa. its elastic modulus is 3.2 gpa. with these properties, it complies with the mentioned standards for static testing. as the iso 14801 standard [12] requires a 30° slanted force, a fixation apparatus is needed to slant the specimens as the litem machine can only apply force in the vertical direction. we needed to make this preparation ourselves as the manufacturer does not provide any further preparations for the machine. it was modelled in netfabb and subsequently 3d printed on the sinterit lisa pro 3d printer by using the sls (selective laser sintering) technology. the used material was the pa12 polyamide, whose strength is comparable with aluminum. the fixating preparation is shown in fig. 7. after the fixation, the implant is connected with the abutment by the abutment screw with a m2.2 thread using a 250 nm momentum by a special ratchet supplied by the manufacturer of application instrumentation. the mechanical tests were controlled by a constant speed of displacement of 4 mm/min until specimen failure in the litem machine. the mean maximum force of the implant specimens at failure was 1327.1 n (table 1) and the failure occured below the head of the abutment screw for all eh specimens, as shown in fig. 5. this test confirmed the hypotheses that the abutment screw would be the weakest part of the assembly and also corrobofigure 5. two eh specimens after mechanical tests. the images show failure of the specimens right below the head of the abutment screw. note that all specimens failed exactly in this manner. figure 6. a sample p-d diagram of specimen n. 2 from the mechanical testing. the maximum force for this specimen was 1387.6 n. rated the research of [6]. from fig. 5, shear failure of all specimens is apparent. the failure of all implants occured by shear failure of the screw connecting the intraosseous part and the abutment. this connection proved to be a limiting factor for the tested eh specimens. a positive finding was the fact that the bodies of the implants themselves remained intact. 2.2. fem analyses as all of the eh specimens achieved failure below the head of the abutment screw, we set out to model these results using ansys workbench 2020 r1. the overall approach was to model the older eh variant, verify the model by comparing it with mechanical tests and subsequently analyse a newer version of the implant with an internal hexagonal connection to confirm whether distributions of stresses are better. the geometrical models were imported into ansys in the form of a parasolid (.x_t extension) and are shown in fig. 9. according to the previously conducted mechanical experiments, we chose a force with the magnitude of 900 n as it is approximately the highest force that is still in the linear region of the p-d diagram (fig. 6). the force is situated on top of the abutment rim and distributed equally among a few nodes of the mesh (fig. 9). to accurately represent the conditions of the experiment and the standard [12], we fixed the lower part of the implant that was embedded in the dentacryl resin and left 83 l. řehounek, a. jíra, g. javorská, d. bodlák acta polytechnica ctu proceedings figure 7. the mechanical testing assembly with the litem machine and a 3d-printed preparation for anchoring the specimens at a 30° angle. a 2 mm gap between the top of the fixed area and the theoretical level of bone (fig. 9). as the goal of the simulation was to compare stress distributions during the linear behavior, we did not consider nonlinear behavior in the model and chose the same force of 900 n for both simulations. simple tetrahedral linear elements and linear material behavior with using the ti-6al-4v material model were used to model the numerical problem. the mesh was designed so as to represent at least 5 elements across the width of the stabilizing ribs of the implant substructure and was further refined by size in the region of interest – the abutment screw. as the assembly consists of 3 solids – the abutment, the screw and the intraosseous part – the model contains basic contact conditions with specific distance break criteria. the contact conditions are necessary as some parts have conical shapes, e.g. the fit is assured by screwing the screw into the abutment conical cavity, the screw passing through the intraosseous part of the implant etc. the contact conditions are shown in fig. 8. due to the complicated geometry of the real screw, we decided to only model its threaded body as a cylinder as the goal of the simulation was not to observe stress distributions in individual threads or details, but rather stress distributions across the whole screw body, i.e. compare the stress concentrations or lack thereof. figure 8. contact regions of the new ih variant inside the ansys environment. figure 9. the two geometrical models inside the ansys environment. left– the old eh variant. right – the new, ih variant. blue shows the area of fixation (2 mm below the theoretical level of bone), red shows the location of the 30° slanted force. boundary conditions comply with previous mechanical tests and with iso 14801 [12]. 84 vol. 30/2021 comparison of transgingival and subgingival implants figure 10. the final results of both analyses (eh and ih variant) of the fem analyses displayed on the abutment screws. a – section through the abutment screw of the eh variant, typical stress concentration marked red. b – a detail of the area highlighted in "a". c – the new ih geometrical solution with significantly reduced shear stress concentrations. 3. results the initial hypothesis was that by introducing a modified geometry of the implant-abutment connection, we can reduce the shear stress concentrations and therefore either eliminate stress concentrations or increase the maximum force at which the screw will eventually break (for the eh variant, the mean force was 1327.1 n, see table 1). the results from the simulation of the old eh variant are shown in fig. 10-a and b. the image shows that the model behaves in the same way as the specimens in mechanical tests. although we did not encompass the region of failure as our analysis is only linear, it is apparent that the concentration of shear stresses occur exactly at the same place – the thin cross-section of the abutment screw below its head. with this information, we constructed the second model in the exact same way and compared the analyses. moreover, another region with significant stress concentration emerged at the bottom of the abutment screw. however, this occurrence was not accounted for as it did not have any direct impact during the mechanical tests. the results from the simulation of the new ih variant are shown in fig. 10-c. the shear stress concentrations were diminished and extremes moved from the thin cross-section of the screw into its head. although the overall values of shear stresses are higher across the whole body of the abutment screw, the concentration points were eliminated by incorporating the internal connection. 4. conclusions the initial hypothesis of this experiment was that by replacing the external hexagonal (eh) connection with internal (ih) connection, we can eliminate stress concentrations that emerge with the use of eh because of its geometry. the mechanical tests have shown that all of the dental implant specimens reached failure during the tests by shear failure of the screw below the its head. von-mises criteria and principal stresses were also considered, but their results were not included in this study as their evaluation has proven to be non-essential for confirming the hypothesis. the numerical analyses have shown that the eh variant induces concentrations of shear stress in the screw below its head. under the same conditions, we 85 l. řehounek, a. jíra, g. javorská, d. bodlák acta polytechnica ctu proceedings performed analysis of the ih variant and found a more viable distribution of shear stress with no apparent concentrations in the body of the abutment screw. therefore, this analysis confirms our initial hypothesis that the internal connection might be more favorable and corroborated the results provided by literature [6]. the trend to innovate in the area of implantabutment connections can be seen in the market share of dental implants, where more and more companies offer alternatives to the tried-and-true external connection, which proves to have its shortcomings, such as low anti-rotational resistance, bone loss at the peri-implant area, lower lateral stability, microgaps at the abutment-implant interface leading to bacteria colonisation and biofilm formation, etc. to further prove if this concept is true and finish the research, we need to manufacture specimens of this new geometrical solution with internal connection and subject them to mechanical tests to determine whether the character of failure changes or the maximum force required to break the abutment screw increases. acknowledgements the financial support provided by the by the faculty of civil engineering, ctu, prague, project n. sgs20/155/ohk1/3t/11 is gratefully acknowledged. references [1] l. gaviria, j. p. salcido, t. guda, j. l. ong. current trends in dental implants. journal of the korean association of oral and maxillofacial surgeons 40(2):50–60, 2014. [2] d. prithviraj, n. muley, v. gupta. the evolution of external and internal implant–abutment connections: a review. international dental research 2(2):37–42, 2012. [3] t. c. truninger, b. stawarczyk, c. r. leutert, et al. bending moments of zirconia and titanium abutments with internal and external implant–abutment connections after aging and chewing simulation. clinical oral implants research 23(1):12–18, 2012. [4] i. sailer, t. sailer, b. stawarczyk, et al. in vitro study of the influence of the type of connection on the fracture load of zirconia abutments with internal and external implant-abutment connections. int j oral maxillofac implants 24(5):850–858, 2009. [5] p. asvanund, s. m. morgano. photoelastic stress analysis of external versus internal implant-abutment connections. the journal of prosthetic dentistry 106(4):266–271, 2011. [6] w. att, s. kurun, t. gerds, j. strub. fracture resistance of single-tooth implant-supported all-ceramic restorations after exposure to the artificial mouth. journal of oral rehabilitation 33(5):380–386, 2006. [7] f. denk, f. denk, a. jíra. a dental implant shaft, 2017. [8] k.-t. koo, e.-j. lee, j.-y. kim, et al. the effect of internal versus external abutment connection modes on crestal bone changes around dental implants: a radiographic analysis. journal of periodontology 83(9):1104–1109, 2012. [9] c. a. a. lemos, f. r. verri, e. a. bonfante, et al. comparison of external and internal implant-abutment connections for implant supported prostheses. a systematic review and meta-analysis. journal of dentistry 70:14–22, 2018. [10] r. caricasulo, l. malchiodi, p. ghensi, et al. the influence of implant-abutment connection to peri-implant bone loss: a systematic review and meta-analysis. clinical implant dentistry and related research 20(4):653–664, 2018. [11] c. alves de sousa, j. j. conforte, k. s. caiaffa, et al. sealing agent reduces formation of single and dualspecies biofilms of candida albicans and enterococcus faecalis on screw joints at the abutment/implant interface. plos one 14(10):e0223148, 2019. [12] s. iof. international standard iso 14801–dentistry–implants-dynamic fatigue test for endosseous dental implants. geneva: international organization for standardization 2007. [13] e. iso. 7500-1. cen en iso 7500-1 metallic materialsverification of static uniaxial testing machines-part 1: tension/compression testing machines-verification and calibration of the force-measuring system. sl: european committee for standardization, august de 2004. document information . 86 acta polytechnica doi:10.14311/app.2020.27.0084 acta polytechnica 27(0):84–89, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app distribution of hydration products in the microstructure of cement pastes michal hlobil czech academy of sciences, institute of theoretical and applied mechanics, prosecká 809/76, 19000 prague, czech republic correspondence: hlobil@itam.cas.cz abstract. this case study focuses on the quantification of the amorphous hydrate distribution in the microstructure of hardened cement paste. microtomographic scans of the hardenend cement paste were thresholded based on histogram image analysis combined with microstructural composition obtained from cemhyd3d hydration model, to separate unhydrated cement grains, crystalline and amorphous hydrates, and capillary pores. the observed spatial distribution of the amorphous hydrate exhibited a strong spatial gradient as the amorphous gel tended to concentrate around dissolving cement grains rather than precipitate uniformly in the available space. a comparative numerical study was carried out to highlight the effect of the spatially (non)uniform hydrate distribution on the compressive strength of the hardened cement paste. keywords: cement paste, hydration products, microstructure, modelling. 1. introduction upon contact with water, the clinker and gypsum grains present in portland cement start to dissolve, releasing ions into the solution. upon saturation, hydration products (i.e. hydrates) begin to precipitate on their surface. with progressing cement dissolution, the hydrates fully cover the grains and form dense reaction rims. these then densify with time, gradually expand into the available pore space, and eventually interconnect each other to form a solid internal network. hydrates in portland cement pastes consist primarily of the amorphous calcium-(aluminate)-silicatehydrate (c-(a)-s-h) gel as well as other crystalline hydrates, namely portlandite (ch), monosulfate (afm), ettringite (aft), and others, depending on the mineral cement composition [1]. the amorphous c-s-h gel occupies the largest volume fraction in the hydrated cement paste microstructure, and is therefore largely responsible for the engineering properties of the material. the distribution of the amorphous gel in space depends on several factors, namely on the particle size distribution of the cement (clinker) used and consequently on its fineness, on the spacing of cement grains in the mixture governed by the water-to-cement ratio, and on the mineralogical cement composition. the microstructure of cementitious composites can nowadays be commonly visualized with a scanning electron microscope (sem), in particular using the backscattered electron mode (bse), provided a suitable specimen cross-section is prepared [2, 3]. in particular, the specimen must be cut, polished and carbon/gold coated for greater viewing sharpness, which adds to a rather laborious preparation [4]. recent developments of in the field of computed microtomography allow to analyze specimens not only without any prior surface treatment, but also allow to see oxides [wt %] st. dev cao 65.04 0.43 sio2 20.60 0.25 al2o3 5.60 0.24 mgo 1.47 0.10 na2o 0.10 0.02 k2o 0.61 0.02 so3 2.31 0.07 p2o5 0.10 0.01 fe2o3 2.49 0.06 table 1. chemical composition of cement id 133 determined by edx. into the specimen, in particular to reconstruct the whole 3d microstructure [5]. this opens the door for the characterization of volume-based characteristics of the microstructure, such as percolation or spatial distribution of phases. 2. materials and methods the cement used in this study was produced in the astm cement and concrete reference laboratory (nist), usa, under the id 133, and can be characterized as a type i/ii ordinary portland cement with a chemical composition listed in table 1. gypsum was added as 5.4 % on a volume basis. the cement is ground to a blaine fineness of about 350 m2/kg and features a cumulative cement particle size distribution as shown in figure 1. fresh cement paste was prepared by mixing conditioned distilled water and cement in a mass ratio of 0.45. subsequently, the paste was poured into molds and kept in saturated conditions until testing. 84 http://dx.doi.org/10.14311/app.2020.27.0084 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 distribution of hydrates in the microstructure of cement pastes 0 0.2 0.4 0.6 0.8 1 0 10 20 30 40 50 60 70 80 bl ai ne fi ne ne ss ≈ 35 0 m 2 /k g w e ig h t fr a c ti o n [ − ] particle size [µm] id 133 figure 1. discretized particle size distribution of cement id 133. 2.1. tomographic scanning of hardened cement paste paste specimens were demolded and scanned using a microtomograph at the european synchrotron radiation facility near grenoble, france, to obtain a three dimensional reconstruction of the cement paste microstructure at selected times. the size of the scanned volume is 1024 × 1024 × 512 volume elements (voxels) with a resolution of 0.95 µm/voxel which provides an effective microstructural volume of 972×972×486 µm3. the signal from the tomograph is converted into a greyscale monochrome image with color intensities proportional to the atomic number of the element encountered in the particular voxel, ranging from full black to bright white similar to the traditional sem/bse imaging. the 3d tomographic image captures the full microstructural features, namely the volumetric distribution of the individual material phases. slicing the full 3d image into individual slices allows to directly visually compare the microstructure with the traditional 2d greyscale sem/bse images. in particular, clinker minerals and gypsum appear as bright areas which are embedded into a mid-tone grey bulk matrix consisting from a mixture of hydration products, and lastly dark grey-to-black areas correspond to capillary porosity, see figure 2. arbitrary segmentation of material phases from the 2d image slices is not straightforward since the 3d volume generally contains a non-negligible noise, and the distinction between hydration products and capillary porosity is challenging. 2.2. modeling of microstructural development since a quantitative microstructural analysis of the investigated hardened cement paste was not available, a simulation of the cement hydration and microstructure development was carried out using the cemhyd3d software package developed by d. bentz at nist, usa [6]. in its third revision [7], cemhyd3d attempts to reconstruct the microstructure of the investigated system by generating a 3d representation figure 2. cropped 2d slice of the tomographic data showing the microstructural features of a hardened cement paste; image size 158 × 158 pixels capturing an area 150 × 150 µm2. of cement particles in water, while matching the particle size distribution of the real cement along with the individual phase volume and area fractions. periodic boundary conditions are respected during the generation of the 3d volume containing cement particles to omit the local segregation phenomena related to the wall-effect, e.g. near the outer specimen surface, or near the aggregates where itz forms [8]. the model accounts for the chemical reactions of the major cement minerals (alite, belite, aluminate, and ferite phases) along with calcium sulfate phases (an-/hemi/di-hydrate). implemented chemical reactions follow the described stoichiometry [9]. computationally, the hydration reactions are implemented as a series of cellular automata-like rules, describing in repeated cycles the dissolution, diffusion, and interaction of diffusing species with clinker and gypsum particles. a posteriori, calibration of hydration kinetics is necessary for mapping the calculated results to real time. this is achieved by fitting the calculated results of the heat evolution into a parabolic dispersion model via time = t0 + β cycle2, (1) where t0 stands for the induction time and β is a fitting parameter, which has to be calibrated from experiments. for the present cement id 133 studied, t0 = 0 as the induction period is implicitly accounted for since the second version of this hydration model [10] by adjusting the early-dissolution probabilities for all clinker phases, which are made proportional to the square of the amount of c-s-h produced, and β = 0.0003 as identified from isocalorimetric measurements carried out in nist. 85 michal hlobil acta polytechnica 0 20 40 60 80 100 0 0.2 0.4 0.6 0.8 1 v o lu m e f ra c ti o n [ % ] degree of clinker hydration ξclin [−] t= 4 .5 h t= 1 1 h t= 1 2 4 h cement crystal. hydrates amorphous hydrate capillary pores figure 3. the volume fraction evolution of microstructural phases in cement id 133, cured at 20 °c in saturated conditions, as simulated by cemhyd3d. figure 4. partially hydrated artificial microstructure from cemhyd3d; one slice showing a cross-section 100 × 100 µm with a resolution 1 pixel (voxel)=1 µm; color map corresponds to figure 3. 3. results the microstructural evolution as quantified by simulating the hydration of an rve of cement paste with a side equal to 100 µm using cemhyd3d is shown in figure 3. to simplify the microstructural characterization, all unhydrated clinker phases together with all calcium sulfate forms (including secondary gypsum formed) are merged together into the phase labeled as “cement”. similar simplification was done also for crystalline hydrates which include portlandite, ettringite, monosulfate, and iron hydroxide. the amorphous hydrate then consists of the calcium silicate and calcium aluminate hydrates. a single 2d slice of the modeled microstructure from cemhyd3d is then shown in figure 4 with a color map matching figure 3. combining histogram image analysis with known (calculated) volume fractions of material phases opens door for accurate thresholding of the tomographic figure 5. thresholded 2d slice of the tomographic data separating cement (clinker+calcium sulfate phases), amorphous and crystalline hydrates, and capillary porosity; image size 158 × 158 pixels capturing an area 150 × 150 µm2; color map corresponds to figure 3. slices, as described next. first, a 2d slice from the tomographic data is extracted and cropped to 158 × 158 pixels capturing an area of 150 × 150 µm to reduce computational time for analysis and yet cover a sufficiently large area for a representative analysis. since the raw tomographic data contains noise, a suitable smoothing image filter is applied to reduce the image noise. subsequently, a histogram of the cropped image is produced, quantifying the amount of pixels of a given grey shade in 8 bit color depth, i.e. 255 colors, ranging from 0 (representing black) to 255 (white). converting the histogram into its cumulative form, and normalizing it with respect to the total amount of pixels in the image allows to keep the abscissa values and renumber the ordinate from 0 to 1. in such form, the cumulative histogram can be used to identify two threshold values for porosity and cement particles by comparing the calculated volume fractions from cemhyd3d with the ordinate values, pointing to the sought threshold values on the abscissa. the greyscale values lying in between these limits are attributed to the hydration products. separation of the hydrate phase into the crystalline and amorphous parts is not possible using histogram analysis, since the phase contrast manifesting as similar grey shade is too low. however, the formation of crystalline hydrates within the microstructure is random as they tend to form both on the surface of cement particles as well as nucleate within the amorphous hydrate [11] consequently, the crystalline hydrates were randomly distributed within the hydration products following their volume fraction, see figure 5 for final result. 86 vol. 27/2020 distribution of hydrates in the microstructure of cement pastes 0 500 1000 1500 2000 2500 3000 0 5 10 15 20 25 30 35 p ix e l c o u n t [− ] minimum distance from cement grain [µm] t=124 h t=11 h t=4.4 h figure 6. quantification of the spatial distribution of amorphous hydrates as a function of the minimum distance from the cement grain. 3.1. spatial distribution of amorphous hydrates quantification of the spatial distribution of the amorphous hydrate gel from the thresholded tomographic image is performed by measuring the distance between the given amorphous hydrate pixel and the closest cement particle. counting such minimum distances for all hydrate pixels in the image and binning the results shows a rather strongly non-uniform hydrate distribution in the cement paste microstructure, see figure 6. in the early stages of hydration (4.5 and 11 hours), the amorphous hydrates form dominantly around cement grains which can be seen as the distance from the grain is rather short, i.e. only single units of micrometers. with progressing hydration, cement grains dissolve and create additional space for new hydrates to form. however, even then the hydrates tend to precipitate in rings, minimizing the distance to the nearest cement grain, and do not precipitate randomly in space. 3.2. prediction of cement paste compressive strength local spatial distribution of hydration products directly affects the macroscopic mechanical properties of cement paste, in particular the compressive strength. two case studies will be considered to show the impact of the hydrate distribution, namely i) the hydrate density distribution exhibiting a strong spatial gradient as observed in thresholded micrographs, see figure 5, and ii) an uniform hydrate distribution. in both cases, the cement grain locations as well as volume fractions of all phases remain identical. the only difference stems from the fact, that in the second case, the hydrates are smeared uniformly throughout the microstructure, which effectively changes the capillary pore size system, compare figures 5 and 7. a micromechanical model based on finite element homogenization is formulated for both cases in such a way that the finite element mesh is generated over the thresholded bitmap micrograph using quadrilateral elements in place of image pixels. local mechanical properties of the identified phases, e.g. of clinker, figure 7. artificially created microstructure based on figure 5, featuring an enforced uniform distribution of hydrates in space; image size 158×158 pixels capturing an area of 150 × 150 µm2; color map corresponds to figure 3. phase e ν ft gf ref [gpa] [–] [mpa] [j/m2] clinker 139.9 0.30 – – [13] amorph. hyd. 24.4 0.24 264.1 4.4 [12] crystal. hyd. 42.3 0.32 – – [14] porosity 0.001 0.49 – – [15] table 2. material properties obtained from nanoindentation. of amorphous and crystalline hydration products are obtained by nanoindentation measurements [12], see table 2. 0 20 40 60 80 100 0 0.002 0.004 0.006 0.008 0.01 0.012 0.014 c o m p re s s iv e s tr e n g th [ m p a ] compressive strain [−] gradient distribution uniform distribution figure 8. results of compressive strength simulation for both hydrate spatial distributions (gradient and uniform) for virtual microstructures of cement id 133 at 124 hrs. two distinct material models are assigned to the individual phases. a simple linear elastic constitutive behavior is used for cement, crystalline hydrates, and 87 michal hlobil acta polytechnica porosity, whereas the amorphous hydrates are modeled using a nonlinear fracture mechanics approach combined with an isotropic damage model for tensile failure, originally described in [16]. the virtual model of the microstructure features periodic boundary conditions in the vertical direction along with supported nodes to prevent rigid body translation and rotation. the problem is solved in the plane stress domain using a nonlinear static analysis carried out with the newton-raphson iterative solver by applying a predefined eigenstrain load increment until failure. results of the microstructural simulation are shown in figure 8. the microstructure based on the thresholded tomographic data exhibits a peak compressive strength of 13.7 mpa, whereas the virtual microstructure exhibiting an uniform distribution of hydrates fails at 78.2 mpa. such an increase of 470% was achieved by changing the hydrate spatial density distribution while keeping the volume fractions of all phases identical. 4. conclusions combining state-of-the-art microtomography to obtain a 3d microstructure with the quantification of volume phases using a hydration model cemhyd3d allowed to view, threshold and separate the cement grains, crystalline and amorphous products, and capillary pores in a representative volume of hardended cement paste with initial water-to-cement ratio of 0.45. focusing on the amorphous hydrate phase, its distribution was measured by determining the minimum distances of each hydrate pixel to the closest cement grain. results of this study indicate a non-uniform distribution with a strong gradient in space. in the early stage of hydration, the hydrates precipitate dominantly around cement grains. with ongoing hydration the distances between cement grains increase as these dissolve, however even then the hydrates tend to minimize the necessary distance and precipitate in rings rather than randomly in space. numerical modeling showed that enforcing an uniform hydrate distribution on the same microstructure (same by means of identical volume fractions of all phases and identical placement of unhydrated cement grains) resulted in a significant increase of compressive strength. this stems from the fact that the uniform hydrate distribution resulted in an alternation of the capillary pore shape and an effective pore size decrease, which consequently manifested as an increase of hardened cement paste compressive strength. acknowledgements the author gratefully acknowledges support from the czech science foundation project 19-25163y. references [1] p. termkhajornkit, q. h. vu, r. barbarulo, et al. dependence of compressive strength on phase assemblage in cement pastes: beyond gel–space ratio experimental evidence and micromechanical modeling. cement and concrete research 56:1–11, 2014. doi:10.1016/j.cemconres.2013.10.007. [2] s. diamond. the microstructure of cement paste and concrete a visual primer. cement and concrete composites 26(8):919–933, 2004. doi:10.1016/j.cemconcomp.2004.02.028. [3] k. scrivener. backscattered electron imaging of cementitious microstructures: understanding and quantification. cement and concrete composites 26(8):935–945, 2004. doi:10.1016/j.cemconcomp.2004.02.029. [4] k. l. scrivener, r. snellings, b. lothenbach. a practical guide to microstructural analysis of cementitious materials. crc press, 2016. [5] e. gallucci, k. scrivener, a. groso, et al. 3d experimental investigation of the microstructure of cement pastes using synchrotron x-ray microtomography (µct). cement and concrete research 2007. [6] d. p. bentz. a three-dimensional cement hydration and microstructure program. i. hydration rate, heat of hydration and chemical shrinkage. tech. rep., nist building and fire research laboratory, gaithersburg, maryland, 1995. [7] d. p. bentz. cemhyd3d: a three-dimensional cement hydration and microstructure development modeling package. version 3.0. tech. rep., nist building and fire research laboratory, gaithersburg, maryland, 2005. [8] k. l. scrivener, a. k. crumbie, p. laugesen. the interfacial transition zone (itz) between cement paste and aggregate in concrete. interface science 12(4):411–421, 2004. [9] h. f. w. taylor. cement chemistry. thomas telford publishing, 2nd edn., 1997. [10] d. p. bentz. cemhyd3d: a three-dimensional cement hydration and microstructure development modeling package. version 2.0. tech. rep., nist building and fire research laboratory, gaithersburg, maryland, 2000. [11] f.-j. ulm, m. vandamme, h. jennings, et al. does microstructure matter for statistical nanoindentation techniques? cement and concrete composites 32(1):92–99, 2010. [12] j. němeček, v. králík, v. šmilauer, et al. tensile strength of hydrated cement paste phases assessed by micro-bending tests and nanoindentation. cement and concrete composites 73:164–173, 2016. doi:10.1016/j.cemconcomp.2016.07.010. [13] k. velez, s. maximilien, d. damidot, et al. determination by nanoindentation of elastic modulus and hardness of pure constituents of portland cement clinker. cement and concrete research 31(4):555–561, 2001. doi:10.1016/s0008-8846(00)00505-6. [14] c.-j. haecker, e. j. garboczi, j. w. bullard, et al. modeling the linear elastic properties of portland cement paste. cement and concrete research 35(10):1948– 1960, 2005. doi:10.1016/j.cemconres.2005.05.001. 88 https://doi.org/10.1016/j.cemconres.2013.10.007 https://doi.org/10.1016/j.cemconcomp.2004.02.028 https://doi.org/10.1016/j.cemconcomp.2004.02.029 https://doi.org/10.1016/j.cemconcomp.2016.07.010 https://doi.org/10.1016/s0008-8846(00)00505-6 https://doi.org/10.1016/j.cemconres.2005.05.001 vol. 27/2020 distribution of hydrates in the microstructure of cement pastes [15] v. šmilauer, z. bittnar. microstructure-based micromechanical prediction of elastic properties in hydrating cement paste. cement and concrete research 36(9):1708–1718, 2006. doi:10.1016/j.cemconres.2006.05.014. [16] m. hlobil, v. šmilauer, g. chanvillard. micromechanical multiscale fracture model for compressive strength of blended cement pastes. cement and concrete research 83:188–202, 2016. 89 https://doi.org/10.1016/j.cemconres.2006.05.014 acta polytechnica 27(0):84–89, 2020 1 introduction 2 materials and methods 2.1 tomographic scanning of hardened cement paste 2.2 modeling of microstructural development 3 results 3.1 spatial distribution of amorphous hydrates 3.2 prediction of cement paste compressive strength 4 conclusions acknowledgements references doi:10.14311/app.2021.30.0126 acta polytechnica ctu proceedings 30:126–130, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app automated bim entity reconstruction from unstructured 3d pointclouds jan voříšek∗, bořek patzák, edita dvořáková, daniel rypl czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jan.vorisek@fsv.cvut.cz abstract. laser scanning is used widely in architecture and construction to document existing buildings by providing accurate data for creating a 3d model. the output is a set of data points in space, so-called point cloud. while point clouds can be directly rendered and inspected, they do not hold any semantics. typically, engineers manually obtain floor plans, structural models, or the whole bim model, which is a very time-consuming task for large building projects. in this contribution, we present the design and concept of a pointcloud2bim library [1]. it provides a set of algorithms for automated or user assisted detection of fundamental entities from scanned point cloud data sets, such as floors, rooms, walls, and openings, and identification of the mutual relationships between them. the entity detection is based on a reasonable degree of human interaction (i.e., expected wall thickness). the results reside in a platform-agnostic json database allowing future integration into any existing bim software. keywords: 3d model, bim, laser scanning, point cloud. 1. introduction the building information modelling (bim) brings excellent opportunities for the efficient and more automated design process of building structures and maintenance and reconstruction planning of existing structures. obtaining a digital bim representation of the latter is often a challenging task, namely when no up-to-date (or none at all) floor plans are available. laser scanning is a convenient and accessible technology to document the existing infrastructure. the output is a dense set of data points in space, socalled point cloud. see figure 1 for an example of a laser-scanned building. each point carries its spatial coordinates, and some scanners also capture the rgb color value and intensity of the scanned point. while point clouds can be directly rendered and inspected, they carry no additional information about the scanned objects. there are no available tools to convert the raw point data into a bim representation with a reasonable degree of automation, although research in this field is highly emerging [2, 3]. currently, the civil engineers have to manually identify the bim entities from graphical visualization of point cloud and create their digital representation in bim model. the long-term goal of the project is to develop an efficient and user friendly tool for semiautomated conversion of point clouds to bim entities while maintaining user interaction. the idea is based on automated identification of entities, where each entity is labeled with estimated degree of identification reliability. the individual identified entities are presented to the user who can accept all suggested entities recognized with reliability higher than defined threshold and manually accepting/modifying remainfigure 1. laser scan of an unfurnished office building, consisting of 85m points, top view. ing suggestions. the presented paper describes the concept of pointcloud2bim library [1] for point cloud manipulation and bim entity identification. 2. pre-processing of the point clouds laser scanners produce very dense point clouds containing tens or hundreds of millions of points. the distance between the scanned points often ranges in millimeters. therefore the point cloud data processing must be handled carefully in order to achieve competitive computation times and memory efficiency. 126 http://dx.doi.org/10.14311/app.2021.30.0126 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 automated bim entity reconstruction figure 2. wide figure. 2.1. storage point information such as the xyz coordinates, rgb color, or intensity value is stored in an uncompressed binary file. the main advantage is that binary data can be directly read into c++ structures instead of the costly parsing of text files line-by-line. all the other project data reside in the project’s json file, which serves as a database of point cloud metadata and other entities. the json format is widely supported across platforms and programming languages and was selected to allow smooth integration into existing software. an empty project file is created automatically before the first point cloud import. 2.2. downsampling while storing the points as binary data offers high reading speeds, with hundreds of millions of points, reading all the values into memory might not be possible on a standard computer with limited memory. for example, storing 100 million xyz coordinates (3x64 bit float) occupies 2,4 gb of memory, which is more than half of a commonly available 4gb. reducing the number of points to work with is inevitable. in our case, the points were downsampled in a 3d voxel grid using the efficient implementation of a c++ hash map [4]. reasonable selection of a grid step of say 1 × 1 × 1 cm reduces all the points within the 1 cm3 cube to a single point with weighted coordinates. result of downsampling to a 1 × 1 × 1 grid is summarized in the table 1. # before after 1 85 316 410 12 358 178 2 62 947 576 20 213 992 3 108 200 686 34 672 667 table 1. the effect of downsampling on three typical point clouds — the number of original points and the number of points after downsampling. one can see that just by downsampling to a 1 cm3 grid, the size of points goes down to at least 30%. 2.3. the normal vector of downsampled points further processing can take advantage of determining the normal vector at each downsampled point (voxel). this is done in the three following steps. a single point represents each voxel by averaging the coordinates of the inner points. there are up to 26 (3 × 3 × 3 − 1) voxels surrounding each voxel if all the neighbors are activated. therefore, the total number of points n for the normal evaluation is at most 27. the centroid of the n points is calculated as a weighted average of the xyz coordinates. µx = !n i=1 xi n , µy = !n i=1 yi n , µz = !n i=1 zi n , (1) the 3 × 3 covariance matrix of the point xyz coordinates relative to the centroid is calculated using the formula for the covariance of a and b below. cov(a, b) = !n i=1(ai − µa)(bi − µb) n − 1 (2) the smallest eigenvector of the covariance matrix is the plane normal [5]. the direct solver of eigenvalues was implemented for the 3 × 3 matrix using the arccos function, as described in [6]. 3. bim entity extraction the overall strategy of point cloud transformation into the digital bim representation of the building is based on a gradual segmentation of floors, rooms, walls, and openings. the overview of this complex process is demonstrated in figure 2. 3.1. floor segmentation the floor segmentation is based on the histogram of vertical coordinates. from the histogram, the ceiling and floor levels are extracted, corresponding to locations with significant location frequency. more accurate results can be obtained by only considering points with vertical normal in the calculation. the ceiling level can also be input manually by inspecting the point cloud in a suitable tool like autodesk recap. 127 j. vorisek, b. patzák, e. dvořáková, d. rypl acta polytechnica ctu proceedings figure 3. results of the room segmentation of the point cloud shown in figure 1. 3.2. room segmentation the ceiling level of the floor is known either from the histogram calculation or from user input. all the ceiling plane points and points within 30 cm below (as suggested in [3]) are projected onto a 2d grid (floor plan) resulting in a pixel-like image. the step of the grid should be at least twice the step used for the initial downsampling, or the grid will contain missing points as some points may not be appropriately captured. region growing [7] is applied to join adjacent pixels into separate regions. backfilling of each of the found regions is performed to include columns and skylights within the room. knowing the pixel count and pixel area (step x step), area of the subject rooms or whole floor can be estimated. finally, the original point cloud is iterated, and points inside of the found regions (in terms of x, y coordinates) and floor (given z coordinate range) are selected. separate point clouds of each room are stored in the project file. an example of room segmentation results can be seen in figure 3. 3.3. vertical plane segmentation the wall planes are identified in each room using the modified random sample consensus (ransac) [8]. only points with horizontal normal are considered for plane finding, a slight deviation of 5 degrees is allowed to account for deviations in construction. each wall finding starts with hypothesis generation. three non-collinear points a, b, c are randomly selected from the point cloud for which the four plane coefficients a, b, c, and d are calculated using the two vectors !ab and !ac !ab = (bx − ax, by − ay , bz − az ) (3) !ac = (cx − ax, cy − ay , cz − az ) (4) figure 4. vertical plane segmentation (top view). door planes are segmented as they cover sufficient area, also notice the noise around outer windows and in the center room. to satisfy the plane equation ax + by + cz + d = 0, plane coefficients are calculated as follows a = (by − ay )(cz − az ) − (cy − ay )(bz − az ) (5) b = (bz − az )(cx − ax) − (cz − az )(bx − ax) (6) c = (bx − ax)(cy − ay ) − (cx − ax)(by − ay ) (7) d = −(aax + bay + caz ) (8) after establishing the hypothesis, more points are selected randomly, and the distance from an arbitrary point to the plane is calculated. points within sufficient distance are considered in the support set. for the first 100 accepted points, the plane coefficients are refined using the algorithm described in section 2.3. 3.4. wall segmentation wall segmentation is performed on each of the vertical planes found in the previous step. only walls of constant thickness are supported. therefore, comparing the angle between every two planes and the distance between the two subject planes is sufficient to mark wall segment candidates. for each wall candidate, its points are projected to a single 2d plane. such a transformation can be done simply without the transformation matrix. the normal vector is taken from plane segmentation by calculating the average value of all wall planes. local x-axis is selected by picking two plane points a, b, and clipping their vector ’s z coordinate. by knowing the local x-axis and the normal vector, the y-axis vector can be calculated using the cross product. this is sumarized below. xloc = (bx − ax, by − ay , 0) (9) yloc = n × xloc (10) 128 vol. 30/2021 automated bim entity reconstruction figure 5. wall segmentation (top view). the wall local axes xloc and yloc are then normalized, and for each wall point p, the local coordinates are calculated using the dot product. px,loc = (p − a) · xloc (11) py,loc = (p − a) · yloc (12) plane points lie in two intervals [a, b] and [c, d] in the local x-axis direction. the distance o in the wall direction in which the planes are overlapping is calculated o = [max(a, c), min(b, d)] wall candidates are filtered based on the user’s overlap requirements, as partially overlapping planes are likely to belong to a common wall. 3.5. opening segmentation a naive way of finding openings in walls (on two parallel planes) is to find overlap of missing points on both planes. however, such treatment produces a lot of false results. for example, two cabinets on each side of the wall in the same place do not form an opening. points present between the walls can be taken into consideration to improve the accuracy of opening detection. by a combination of inverse region growing and boundary detection, openings with a sufficient density of intermediate points can be identified. by calculating the bounding box of an opening in the local plane coordinates, dimensions can be calculated for rectangular openings (i.e., doors and some windows). doors are recognized by comparing their bottom coordinates to the corresponding floor level. opening visualized by their bounding box can be observed in figure 6. figure 6. opening detection — doors (red), windows (blue). 4. demonstration on a sample point cloud the example of a simple office building with just a single floor with eight rooms is selected to demonstrate the capabilities of the presented approach. as the whole floor was unfurnished, the overall floor envelope has little to none missing points except for the areas behind doors and heating elements. the original data set consists of 85 316 410 points is shown in fig. 1. at first, the original point cloud was downsampled into 12 358 178 using a 1 × 1 × 1 cm grid. by inspecting the point cloud, we identify the floor and ceiling levels, and we use all the points in between ceiling level and 30 cm below for room finding. the room finding performed on a 2d floor plan grid with a 2 × 2 cm step yields all the eight identified rooms, visualized in figure 3. in the same step, we also identify all the 93 vertical planes, as can be seen in fig. 4. the planes assigned to their source rooms are then passed to the next step for wall matching even if some of the identified planes might be out of our interest (i.e., doors or noise). in figure 5, we can observe all the matched walls on the subject floor. while all the inner walls were identified, the exterior walls are missing as the exterior scan of the office building was not available. however, some of the external walls are obtained thanks to the scan capturing the window frames’ internal and external planes. these reveal the estimated window positions as well as exterior wall location, excluding its wall thickness. 5. conclusions the library was implemented in c++ based on the previously formulated methodology. it can be inte129 j. vorisek, b. patzák, e. dvořáková, d. rypl acta polytechnica ctu proceedings grated into any bim software to assist the conversion of point clouds into digital bim representation. the library api is documented in the reference manual [9]. the library also works as a standalone toolkit without any dependency on bim software. a set of command-line executables is available to perform the individual steps introduced above. each of the tools operates on top of a common database-like file (json) with point data stored separately in binary files. a single floor of an unfurnished office building was successfully segmented into separate rooms (compare figures 1 and 3). plane matching results can be seen in figure 4 with the 25 distinct wall segments considering the exterior windows as walls. the success of opening detection can be examined in figure 6 where six doors were identified and eight exterior windows (two of which are found twice). 6. future work our main goal is to create a robust and easy-tointegrate solution of converting the raw point cloud files into the corresponding bim entities. while our library provides enough functionality to process, analyze, and visualize the point clouds and calculated results, more research is desired, especially in the area of opening detection and overall confidence rate evaluation of the reconstructed bim entities. further research is needed to integrate the library into some of the industry-leading bim applications (i.e. autodesk revit). such integration is required to validate and refine our methodology using a variety of point clouds from different environments. acknowledgements the development of this software was supported by the technology agency of the czech republic under the program of the national competence center 1 as a project “center for advanced materials and efficient buildings” (cameb) project registration no. tn01000056 and by the grant agency of the czech technical university in prague (sgs project no. sgs20/038/ohk1/1t/11), both gratefully acknowledged. references [1] pointcloud2bim library for automated identification of bim entities from point clouds. http://mech2018.fsv.cvut.cz/pc2bim. accessed: 2020-08-06. 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[7] m. montoya, c. gil, i. garcía fernandez. implementation of a region growing algorithm on multicomputers: analysis of the work load balance., 2000. [8] m. a. fischler, r. c. bolles. random sample consensus: a paradigm for model fitting with applications to image analysis and automated cartography. communications of the acm 24(6):381–395, 1981. doi:10.1145/358669.358692. [9] j. voříšek, d. rypl, b. patzák, e. dvořáková. pointcloud2bim library: pointcloud2bim library reference manual. http://mech2018.fsv.cvut.cz/ pc2bim/www/refman/html/index.html. 130 http://mech2018.fsv.cvut.cz/pc2bim https://doi.org/10.1016/j.autcon.2015.04.001 https://doi.org/10.3390/app7101030 https://greg7mdp.github.io/parallel-hashmap/ https://www.ilikebigbits.com/2017_09_25_plane_from_points_2.html https://doi.org/10.1145/358669.358692 http://mech2018.fsv.cvut.cz/pc2bim/www/refman/html/index.html acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0039 acta polytechnica ctu proceedings 26:1–6, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app modelling of crack propagation: comparison of discrete lattice system and cohesive zone model karel mikeša, ∗, franz bormannb, ondřej rokoša, b, ron h.j. peerlingsb a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b eindhoven university of technology, department of mechanical engineering, p.o. box 513, 5600 mb, eindhoven, the netherlands ∗ corresponding author: karel.mikes@cvut.cz abstract. lattice models are often used to analyze materials with discrete micro-structures mainly due to their ability to accurately reflect behaviour of individual fibres or struts and capture macroscopic phenomena such as crack initiation, propagation, or branching. due to the excessive number of discrete interactions, however, such models are often computationally expensive or even intractable for realistic problem dimensions. simplifications therefore need to be adopted, which allow for efficient yet accurate modelling of engineering applications. for crack propagation modelling, the underlying discrete microstructure is typically replaced with an effective continuum, whereas the crack is inserted as an infinitely thin cohesive zone with a specific traction-separation law. in this work, the accuracy and efficiency of such an effective cohesive zone model is evaluated against the full lattice representation for an example of crack propagation in a three-point bending test. the variational formulation of both models is provided, and obtained results are compared for brittle and ductile behaviour of the underlying lattice in terms of force-displacement curves, crack opening diagrams, and crack length evolutions. the influence of the thickness of the process zone, which is present in the full lattice model but neglected in the effective cohesive zone model, is studied in detail. keywords: lattice model, damage, finite element method, cohesive zone model, three-point-bending test, crack propagation. 1. introduction lattice models with dissipative interactions are often used for modelling of materials with discrete microstructures such as textile, paper, 3d printed materials, or concrete; see e.g. [1–3]. the advantage of lattice models lies in their ability to capture microstructural interactions individually, thus representing the geometry of the underlying micro-structure as well as the associated constitutive laws accurately and naturally. because non-local behaviour is important for mechanisms occurring in the process zone at the crack tip, discrete models are suitable for problems such as damage localization, crack propagation, and branching. lattice models suffer, on the other hand, from high computational and memory requirements, necessitating simplifications especially in cases in which multiple evaluations are required, such as optimization, material parameter identification, or stochastic simulations. although multiple options for building effective models exist, in this contribution we focus on a continuum approximation of the underlying microstructure with an embedded cohesive zone model. crack propagation and its cohesive properties are described in such an approach by a suitable traction-separation law, whereas the bulk behavior of the continuum may often be elastic. the aim of this paper is detail comparison of crack propagation for discrete lattice model and continuous cohesive zone approach. the accuracy and performance of cohesive zone model predictions are assessed on the example of crack propagation in a three-pointbending test, as shown in figure 1, by comparison with the exact underlying full lattice model. 2. full lattice model an admissible configuration of a rate-independent system of interest is specified by a state variable q = (r,z), where r stores the kinematic variables (displacement of nodes) and z the internal variables (damage in bonds). the energetic solution q is determined using the following incremental minimization problem at time tk q(tk) ∈ arg min q̂∈q πk(q̂; q(tk−1)), k = 1, . . . ,nt , (1) with an initial condition q(0) = q0, where q denotes the space of all admissible configurations, and πk is an incremental energy defined as πk(q; q(tk−1)) = e(tk,q) + d(z,z(tk−1)). (2) the incremental energy consists of the potential (gibbs type) energy e and the dissipation distance d. 39 https://doi.org/10.14311/app.2020.26.0039 https://ojs.cvut.cz/ojs/index.php/app k. mikeš, f. bormann, o. rokoš, r.h.j. peerlings acta polytechnica ctu proceedings l h f ~ex ~ey ωa figure 1. geometry and boundary conditions of the considered three-point-bending test with a close-up of the full lattice representation. the potential energy reads e(tk,q) = v(q) −ftext(tk)r, (3) where v is the internal free energy and fext is a vector of external loads. the dissipation distance d(z2,z1) measures the minimum dissipation by a continuous transition between two consecutive states z1 and z2; for further details see [4] (section 3.2). it can be written as d(z2,z1) = 1 2 ∑ α,β∈bα dαβ(z2,z1), (4) where bα denotes the set of nodes connected with node α. dαβ is the dissipation distance for a single bond αβ, defined as follows dαβ(z2,z1) = { dαβ(ω2) −dαβ(ω1) if ω2 ≥ ω1 +∞ otherwise, (5) where dαβ(ω) reflects the amount of energy dissipated in that bond during a unidirectional damage process up to a given damage level ω. in each time step, the minimization problem of eq. (1) is solved to obtain a local minimum that satisfies the energy balance v(q(tk)) −v(q(0)) + vard(q, tk) = wext(q, tk), (6) which equates the internally stored energy v(q) − v(q0) plus the dissipated energy vard(q, tk) = k∑ i=1 d(z(ti),z(ti−1)), (7) with the work performed by the external forces wext(q, tk) = k∑ i=1 1 2 [f(ti) + f(ti−1)]t[r(ti) −r(ti−1)]. (8) unless stated otherwise, the same material model is considered for each lattice bond. in undamaged state, it is defined by pair potential φαβ(lαβ) = 1 2 e0a0l αβ 0 ( ε(lαβ) )2 , (9) where e0 is the young’s modulus and a0 the crosssectional area of the bond. l0 denotes the initial length parameter ε0 εf e0 a0 brittle 0.005 10ε0 1/ε0 1 ductile 0.005 100ε0 1/ε0 1 table 1. parameters of the lattice model. of the bond and ε(lαβ) = (lαβ − l0)/l0 is the bond strain. the internal energy is defined as v(r,z) = 1 2 ∑ α,β∈bα [ (1 −ωαβ)φαβ(lαβ) ] . (10) the elastic energy of a bond φαβ is reduced by the damage variable ωαβ(ε), defined as ωαβ(ε) = { 1 −ε0 exp ( −ε−ε0 εf ) if ε ≥ ε0 0 if ε < ε0, (11) where ε0 is the limit elastic strain, and the parameter εf characterizes the softening branch of the associated stress-strain diagram. the particular choice of material and geometric parameters corresponding to the lattice model are specified in table 1. for large lattice structures, the simulation must be controlled by the increment of dissipation, cf. e.g. [5], because both the (typically used) crack mouth opening displacement control (cmod) as well as crack tip displacement opening control (ctod) turned out to be insufficient to keep the energy balance of eq. (6) satisfied. 3. cohesive zone model the lattice problem as specified in figure 1 is approximated with an effective continuous finite element cohesive zone model (fe-cz). for that purpose, the following assumptions are adopted: (1.) only one crack is initiated; (2.) the crack trajectory is a straight line and the crack propagates vertically in the middle of the beam; (3.) all damage processes take place in the predefined crack path, whereas the rest of the domain remains elastic. 40 vol. 26/2020 comparison of discrete lattice system and cohesive zone model ωar (elastic region) l h f γcz ωal (elastic region) cohesive zone with traction separation law figure 2. geometry and boundary conditions of the finite element cohesive zone model (fe-cz) subjected to a three-point bending test: elastic domain (grey) and cohesive zone (dashed line). 3.1. model description the crack is represented as one cohesive zone, γcz, normal to the ~ex direction, which splits the considered domain ω into two symmetric parts ωl and ωr: γcz = { ~x ∈ r2 : x = xcz = 0, |y| ≤ h/2 } , ωl = { ~x ∈ r2 : −l/2 ≤ x < xcz, |y| ≤ h/2 } , ωr = { ~x ∈ r2 : xcz < x ≤ l/2, |y| ≤ h/2 } , ω = ωl ∪ ωr, (12) where xcz = 0 denotes the horizontal position of the vertical cohesive zone, cf. figure 2. the total free energy ψ of the fe-cz model consists of two contributions, i.e. ψ = ∫ ω\γcz ψe dω + ∫ γcz ψcz dγcz, (13) where ψe is the elastic strain energy density ψe = 1 2 ε : c : ε, (14) corresponding to a linear elastic isotropic material under plane strain assumptions, described through a fourth-order constitutive tensor c, even though the full lattice is not isotropic. in eq. (14), the elastic strain tensor is expressed as the symmetric part of the displacement gradient ~∇~u, i.e. ε = 1 2 ( ~∇~u + ( ~∇~u )t) . (15) the cohesive zone potential ψcz is a function of the displacement jump ~∆ = j~uk = ~ur −~ul (16) between the two sub-domains ωr and ωl. due to the symmetry of the problem, the tangential component of the jump ∆t vanishes along the cohesive interface, i.e. ~uly = ~u r y , on γcz, (17) meaning that the cohesive zone potential ψcz becomes a function of the normal crack opening ∆n only. the constitutive parameters of both models, i.e. c and ψcz(∆n), are fitted in section 3.2. for the solution, in each loading step the resulting total potential energy of eq. (13) is minimized with respect to the degrees of freedom associated with a discretized displacement field of the continuum model. in order to handle brittle behaviour and snap-back, cmod indirect control is used. 3.2. material parameter calibration both effective models, i.e. the cohesive zone potential ψcz and elasticity stiffness tensor c, are identified directly from the properties of the underlying lattice by homogenization. for that purpose, a single periodic unit cell is considered, as shown in fig. 3. note that the cross-sectional areas of the horizontal and vertical bounding links are reduced by a factor of 0.5 to account for periodicity. l0 l0 ∆n ∆n figure 3. the geometry of unit lattice used for calibration of elastic properties and traction-separation law. by homogenization, the effective stiffness tensor can be assembled and compared with the plane strain formula for hooke’s law to obtain the elastic coefficients. in particular, the effective young’s modulus reads eiso = 5 6 e0a0 l0 (1 + √ 2/2), (18) whereas poisson’s ratio ν is fixed at ν = 1/4 for the given lattice geometry; for more details see e.g. [6, appendix a]. 41 k. mikeš, f. bormann, o. rokoš, r.h.j. peerlings acta polytechnica ctu proceedings figure 4. normalized traction as a function of normalized opening for the two parameter sets considered, cf. tab. 1. figure 5. the damage (red) in bonds computed with full lattice model: evolution of the process zone one step before crack initiation (top) and final process zone and fully developed crack path in the last step of simulation (bottom). instead of the cohesive zone potential ψcz, the corresponding traction, t , defined as t(∆n) = dψcz(∆n) d∆n , (19) is evaluated numerically from the response of the periodic unit cell subjected to a prescribed displacement ∆n with the boundary conditions according to figure 3. the computed values of tractions corresponding to the two different parameter sets of tab. 1 are plotted in figure 4. the cohesive zone potential ψcz can be obtained from these tractions by numerical integration. but effectively, for the practical implementation, the values of ψcz are not needed because the problem is solved for the force balance by the standard newton algorithm. 4. results in this section, the full lattice model defined in section 2 (referred to as full lattice), and finite element cohesive zone model defined in section 3 (referred to as fe-cz ), are compared on a three-point-bending test of a rectangular specimen of dimensions 2048l0×256l0, for the brittle and ductile lattice material settings (recall table 1 and figure 4). for the full lattice model, individual bonds start to damage in a relatively ductile manner, resulting in a considerable process zone, featuring multiple possible crack paths, cf. figure 5 (top). after localization, a full (central) crack develops across the entire height of the specimen, cf. figure 5 (bottom). the force-displacement responses are compared in figure 6. the fe-cz is able to exactly capture the initial elastic stiffness. however, the peak force is significantly underestimated by fe-cz , for both materials. the relative peak force error (of fe-cz in comparison with full lattice) is 23.3 % for the brittle and even 33.4 % for the ductile material. this is caused by the fact that for the full lattice a large amount of inelastic strain along the bottom edge is accumulated in the process zone. this complex behavior cannot be captured with one cohesive zone and otherwise elastic behavior naturally results in significant deviations in the fe-cz model. the inelastic process zone in the full lattice unloads the crack mouth (compared to fe-cz ) and, therefore, higher loading force is needed to initiate the crack. to verify this, we limit the process zone in the lattice model in a further simulation (referred to as elastic lattice). in this model, the damage can evolve only in a narrow band of width 2l0 positioned along the vertical axis of the symmetry, whereas the remaining lattice is considered as purely elastic. to this end, the critical strain parameter is modified as εi0 = { ε0 if |xi −xcz| ≤ l0 ∞ othervise (20) 42 vol. 26/2020 comparison of discrete lattice system and cohesive zone model figure 6. normalized force–displacement diagrams for a brittle (left) and a ductile (right) lattice. the results of elastic lattice correspond to the lattice where the damage can evolve only in a narrow band of width 2l0, whereas the remaining lattice is elastic. figure 7. normalized opening profiles for fixed pre-peak loading force f/(e0a0ε0) = 56 (left) and for post-peak loading force f/(e0a0ε0) = 14 (right) computed for the ductile material setting. figure 8. normalized crack length as a function of crack mouth opening displacement (cmod) computed for the ductile material setting. where xi −xcz is the horizontal distance between the center of the i-th bond and the symmetry axis. for all considered aspects, the fe-cz model provides very similar results to the elastic lattice, where the influence of the process zone is removed. it reveals that the cohesive zone in fe-cz is able to exactly capture the damage properties of the underlying lattice and the error in the peak force is produced mainly by the missing process zone. this can also be observed in the difference of preand post-peak opening profiles, corresponding to a fixed loading force, as shown in figure 7 for two magnitudes of the applied force. here we clearly see that the fe-cz and elastic lattice (both with a narrow process zone) tend to show significantly larger crack mouth openings before crack localization. once the crack localizes and opens, however, all models provide very accurate results. also the kinematics of crack propagation, depicted in figure 8, is accurately captured by both the elastic lattice and fe-cz . the fe-cz (compared to the full lattice model in computing time) provides, however, a substantial speed-up of the order of 50. 5. conclusion in this contribution, a cohesive-zone finite element approach to modelling of crack propagation in materials with discrete microstructures has been compared against the full lattice model that features phenomena observed in real materials, such as non-local behaviour, large deformations and rotations of individual struts, and distributed cracking. it has been shown that the cohesive-zone model achieves a substantial speed up, while being sufficiently accurate in terms of kinematic quantities such as overall displacement, crack length, or crack opening profile. on the other hand, the conjugate quantities, 43 k. mikeš, f. bormann, o. rokoš, r.h.j. peerlings acta polytechnica ctu proceedings such as the peak force, are significantly underestimated due to the missing process zone which results in a significantly more ductile response of the considered lattice system. with increasing ductility of the lattice, the accuracy further decreases. the presented results have also shown that in order to correct for this deficiency, more ductile behaviour should be accounted for, e.g., by adding multiple (parallel) cohesive zones that cover the size of the main part of the process zone. however, the spatial positioning and orientation of such secondary cohesive zones is not known and needs to be based on prior knowledge. in the most general case, every inter-element interface can be considered as a cohesive zone, which might significantly complicate the entire procedure, introduce a dependence on the discretization (finite element size and orientation), and cause significant computational overhead. let us note that, apart from the homogenization procedures used in this paper, the effective cohesive-zone constitutive parameters can be obtained by fitting directly the overall macroscopic force–displacement curves. although this might improve the macroscopic accuracy (in terms of force–displacement curve) achieved by the cohesive-zone model, such a phenomenological approach would not directly represent the underlying microstructure. this would in turn mean that different geometries would require independent identifications and that the accuracy in kinematic quantities (such as crack opening profile) might be decreased. the cohesive zone model provides an efficient and effective tool for the description of crack propagation in discrete systems, although care must be taken to position the cohesive zone interface properly to cover not only the crack trajectories, but also phenomena important for given problem. list of symbols q state variable r kinematic variable z internal variable t parametrization time nt number of time steps q space of all admissible configurations πk incremental energy e potential energy d dissipation distance v internal free energy vard dissipated energy fext vector of external loads α,β node indices bα nearest-neighbour nodes to α ωαβ damage level in a bond αβ dαβ energy dissipated in bond αβ wext work performed by the external forces φ pair potential e0 young’s modulus of individual bonds a0 cross-sectional area l length of deformed bond l0 initial lattice spacing ε,ε0 bond strain, critical bond strain εf damage softening parameter ω elastic domain γcz cohesive zone ψ total free energy ψcz cohesive zone potential ψe elastic strain energy density ε elastic strain tensor σ elastic stress tensor c elastic stiffness tensor ~u displacement field ∆n, ∆t normal/tangential jump in displacements eiso effective young’s modulus ν effective poison’s ratio f magnitude of the point loading force u displacement under the point loading force x,y cartesian coordinates acknowledgements financial support of this work from the grant agency of the czech technical university in prague (sgs) under project no. 19/032/ohk1/1t/11 and from the czech science foundation (gačr) under project no. 17-04150j is gratefully acknowledged. references [1] l. beex, c. verberne, r. peerlings. experimental identification of a lattice model for woven fabrics: application to electronic textile. composites part a: applied science and manufacturing 48:82–92, 2013. doi:10.1016/j.compositesa.2012.12.014. [2] e. bosco, r. h. peerlings, m. g. geers. predicting hygro-elastic properties of paper sheets based on an idealized model of the underlying fibrous network. international journal of solids and structures 56:43–52, 2015. doi:10.1016/j.ijsolstr.2014.12.006. [3] g. cusatis, z. p. bažant, l. cedolin. confinement-shear lattice model for concrete damage in tension and compression: i. theory. journal of engineering mechanics 129(12):1439–1448, 2003. doi:10.1061/(asce)0733-9399(2003)129:12(1439). [4] a. mielke, t. roubíček. rate-independent systems: theory and application. springer-verlag new york, 1st edn., 2015. doi:10.1007/978-1-4939-2706-7. [5] c. v. verhoosel, j. j. remmers, m. a. gutiérrez. a dissipation-based arc-length method for robust simulation of brittle and ductile failure. international journal for numerical methods in engineering 77(9):1290–1321, 2009. doi:10.1002/nme.2447. [6] k. mikeš, m. jirásek. quasicontinuum method extended to irregular lattices. computers & structures 192:50–70, 2017. doi:10.1016/j.compstruc.2017.07.002. 44 https://doi.org/10.1016/j.compositesa.2012.12.014 https://doi.org/10.1016/j.ijsolstr.2014.12.006 https://doi.org/10.1061/(asce)0733-9399(2003)129:12(1439) https://doi.org/10.1007/978-1-4939-2706-7 https://doi.org/10.1002/nme.2447 https://doi.org/10.1016/j.compstruc.2017.07.002 acta polytechnica ctu proceedings 26:1–6, 2020 1 introduction 2 full lattice model 3 cohesive zone model 3.1 model description 3.2 material parameter calibration 4 results 5 conclusion list of symbols acknowledgements references https://doi.org/10.14311/app.2022.33.0533 acta polytechnica ctu proceedings 33:533–538, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague improvement mechanism of recycled concrete using c-s-h type accelerator and modifying recycled aggregate by carbonation technology akari shibuyaa, ∗, nobuhiro matsudab, takeshi iyodac a former graduate school of engineering and science, shibaura institute of technology, tokyo, japan b tokyo techno co., ltd., tokyo, japan c shibaura institute of technology , tokyo, japan ∗ corresponding author: me19047@shibaura-it.ac.jp abstract. in japan, there is concern that concrete blocks due to demolition will increase with the renewal of concrete structures. although it has been used for roadbed materials as recycling, in recent years, aggregates are removed from the concrete mass by crushing and grinding and used again as concrete aggregates. in concrete using recycled aggregates, it is necessary to take account of insufficient strength, drying shrinkage, freeze-thaw resistance. in previous study, the quality of recycled aggregates was improved by adsorbing carbon dioxide to recycled aggregates, and recycled concrete used the modified aggregates has been confirmed to increase strength and reduce drying shrinkage. in this study, we focused on the aggregate interface of concrete such as the interfacial transition zone for the purpose of clearing the quality improvement method of recycled concrete. we attempted to modify the aggregate interface using a c-s-h hardening accelerator using two recycled aggregates of different quality and concrete using ordinary aggregates. we will also compare which is more effective, concrete that tried to modify the aggregate interface and recycled concrete that uses carbonated modified recycled aggregate. we try include strength test, drying shrinkage test, and freeze-thaw test. we clarify the matters concerned about recycled concrete. keywords: c-s-h hardening accelerator, mass transfer resistance, recycled aggregate, recycled concrete. 1. introduction in japan, structures built during the period of high economic growth are increasingly required to be rebuilt or repaired over their life. 30 million tons of concrete are demolished annually. demolition concrete are increasing year by year. in addition, the readymixed concrete left over and returned to the factory is called the return concrete, and the amount of return concrete is increasing. recently, recycled aggregate is attracting attention from the viewpoint of reducing environmental impact. the recycled aggregate is an aggregate obtained by removing the aggregate from demolition concrete or return concrete so that it can be reused. the quality of recycled aggregate varies depending on the condition and processing method of the original concrete demolition material. in japan, there are three levels of quality, l, m and h. standards of recycled aggregate and concrete are specified in jis a 5021 ∼ 5023. the worst quality l is aggregate that has just been crushed and has a lot of mortar attached. on the other hand, l has merits such as lower energy cost than others. it was thought that low-quality recycled aggregate could be modified. various problems due to the large amount of attached mortar. the big problem is the decrease in strength due to many pores. in addition, high water absorption results in large drying shrinkage, causing freeze-thawing and scaling problems. therefore, the purpose of this study is to understand the problems of recycled aggregate and to examine methods that can solve these problems. in previous studies, carbonation of aggregates has been reported as a method of modifying recycled aggregates. it is reported that strength, drying shrinkage, and freezethaw resistance were improved. by forcibly carbonating the aggregate, calcium carbonate is generated in the attached mortar. pores decrease and mortar area becomes denser. on the other hand, previous studies have reported that calcium silicate hydrate type accelerator (acx) is effective in densifying the interfaced at the aggregate interface [1]. acx is a curing accelerator developed in recent years. it is mainly composed of nanoparticles of calcium silicate hydrate. it is reported that adding 5% of acx to water densifies the voids in the transition zone around the aggregate. by using acx, which is said to be effective at the aggregate interface, it is thought it can reform recycled concrete using recycled aggregate. in this study, we try to understand their impact on concrete by comparing recycled concrete with carbonation of aggregate and recycled concrete with acx added. 533 https://doi.org/10.14311/app.2022.33.0533 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en a. shibuya, n. matsuda, t. iyoda acta polytechnica ctu proceedings surface dry density! g/cm3 " dry density! g/cm3 " water absorption [%] rl 2.42 2.29 5.53 rc 2.46 2.36 4.38 table 1. physical properties of coarse aggregate. 5;:� � � � � � � � � � � � 5;6� figure 1. overview of rxw and rxs. w/b [%] s/a [%] air [%] quantity of material per unit volume of concrete acx [w × %] fresh properties w opc ggbfs s g air[%] slump [cm] r 50 48 4.5 170 187 153 847 rl847 0 3.5 22.5 rx3 3 3.9 22.2 rx10 10 2.4 22.8 rxw 1* 4.1 23.1 rxs 2* 2.3 23.0 rc rl861 0 4.2 23.4 1* : pre-wetting with 10% strength acx 2* : spraying the same amount of acx as in rx3 table 2. mix proportion. 2. materials and specimen condition 2.1. the kinds of coarse aggregates it was used recycled aggregate (rl). in addition, the recycled aggregate is carbonated in an environment with a carbon dioxide concentration of 5% (rc). table 1 shows their physical properties. rl has high water absorption, and carbonation has the effect of reducing water absorption. before mixing, rl was pre-soaked for 24 hours and the surface dried because the recycled aggregates have high water absorption. 2.2. mix proportions table 2 shows the mix proportions on this research. concrete with a water binder ratio of 50% was prepared by replacing the ground granulated blast furnace slag by 45%. concrete using the recycled aggregate was similarly prepared (r, rx3, rx10) and in addition, the concrete using the two types aggregate. one is the pre-wetting was pre-soaked with the acx solution of 10% concentration for 24 hours (rxw) and the other is the concrete that acx was sprayed over all sides of the aggregate (rxs) as shown in figure 1. 3. outline of experiment 3.1. bleeding test bleeding was tested to determine the effect of bleeding suppression by the addition of acx and the modification of aggregate in accordance with jis a 1123. concrete was cast into formwork that does not penetrate water, and the formwork size is φ100 × 200 mm. the test was started by lowering the casting surface by 30mm ± 3mm. every 10 minutes after the casting, the water that had exuded on the upper surface was sucked with a dropper and transferred to a graduated cylinder for measurement. measurements were taken until no bleeding was observed. the bleeding amount was determined by equation 1. 534 vol. 33/2022 cdw structural concrete figure 2. air permeability test. figure 3. bleeding amount. bq = v a , (1) where • bq : amount of bleeding, • v : total bleeding water volume until the end • a: area of concrete casting surface. 3.2. drying shrinkage test two pieces of concrete of 100×100×400 mm with pins were prepared. the measured value after 7 days of underwater curing was used as the initial value, and after curing, the specimen stands in a constant temperature and humidity room (20◦c, 60%rh). the length was measured after 1,4,14, and 28 days in accordance with jis a 1129. 3.3. compressive strength test and splitting tensile strength test three specimens of φ100 × 200 mm were prepared for each case, and after curing in water for 28 days, a compressive strength was tested in accordance with jis a 1108. also, a splitting tensile strength test was performed in accordance with jis a 1113. 3.4. air permeability test specimens of 150 × 150 × 150 mm was prepared and cured for 28 days. after extracting the core to the size of and leaving it in a completely dry state, an air permeability test was performed. in order to understand the effects of bleeding and the quality of the aggregate interface, air permeability test was performed from two directions as shown in figure 2. in concrete, water rises to the upper surface, so it is considered that there are places where the water rises, and where there is a void after hardening due to water that could not rise and accumulated on the lower surface of the aggregate. the air permeability coefficient was calculated by using equation 2. k = 2 l p1 (p 21 − p 22 ) · q a , (2) where • k: air permeability coefficient [cm4/(n·s)] • l: sample thickness [cm] • p1: loading pressure ([n/cm2] • p2: outflow side pressure [n/cm2] • q: air permeability [cm3/s] • a: air permeability area [cm2] 3.5. porosity test the porosity was calculated by the archimedes method. the weight saturated with water (w1) and the weight in water (w2) were measured, and then the specimens stand in a chamber at 40◦c. until the weight loss became constant, and the dry weight (w3) was measured. the weight was calculated using the mass, and the weight was calculated by equation 3. 535 a. shibuya, n. matsuda, t. iyoda acta polytechnica ctu proceedings porosity [%] = w1 − w3 w1 − w2 × 100. (3) figure 4. drying shrinkage rate. figure 5. compressive strength and splitting tensile strength. 4. results 4.1. bleeding test in figure 3, the total amount of bleeding is organized for each specimen. rx3 and rx10 showed non-bleeding. it became clear that the addition of acx to concrete has a large inhibitory effect. rxw and rxs almost showed inhibitory effect. the reason that rxs had a smaller bleeding amount than rxw is probably that acx sprayed on the surface diffused during mixing and had the same effect as that added. 4.2. drying shrinkage test figure 4 shows the results of drying shrinkage at 28 days. drying shrinkage was suppressed by adding acx. it became clear that adding 10% was more effective than adding 3%. also, rxw, rxs and rc has almost same shrinkage rate. they did not show a reduction effect more than when acx was added to the concrete itself. therefore, drying shrinkage is best when acx is added to the concrete mixing water. 4.3. results of strength test for concrete using recycled aggregate, rx10 had higher compressive strength, however rxw and rxs did not get stronger. therefore, adding acx to concrete is effective for compressive strength, however figure 6. air permeability coefficient (horizontal). figure 7. compressive strength and air permeability coefficient. figure 8. air permeability coefficient in horizontal and vertical direction. methods such as pre-wetting and spraying that attempt to modify the aggregate itself have no effect. the greater the amount of acx added, the greater the splitting tensile strength. rxw showed no increase in strength, but rxs showed the same degree of increase in strength as the addition of acx. a possible reason for this is that the pre-wetting aggregate was not able to play a role in modifying the aggregate interface by wiping the surface. in addition, rc showed the highest strength among concrete using recycled aggregate. comparing the compressive strength with the split tensile strength, the effect on the split tensile strength is seen in rc. therefore, it was considered that the splitting tensile strength increased because the strength of the aggregate itself increased due to carbonation, but the compressive strength of the aggre536 vol. 33/2022 cdw structural concrete �� �� �� �� �� �� �� �� & rp su hv vl yh �v wu hq jw k� �1 �p p � � 3rurvlw\���� 5 5;� 5;�� 5;: 5;6 5& ��� ��� ��� ��� � �� �� �� 6s olw wlq j� wh qv loh �v wu hq jw k� �1 �p p � � �� 3rurvlw\���� 5 5;� 5;�� 5;: 5;6 5& figure 9. the relationship between porosity and strength. figure 10. the relationship between porosity and air permeability. gate interface and mortar did not change. 4.4. air permeability test 4.4.1. air permeability coefficient figure 6 shows the permeability coefficient calculated from the result of the permeability test in the horizontal direction on recycled concrete. the significant improvement of the air permeability by adding acx is apparent. due to the modification of the recycled aggregate itself, it is considered that rxw and rc had been densified by the mortarattached to the recycled aggregate. 4.4.2. the relationship between compressive strength and air permeability coefficient as shown in figure 7, for concretes other than rxs, there was a correlation that the compressive strength decreased as the air permeability decreased. it was cleared that rx10 had significantly improved concrete. only the permeability was improved with almost the same strength excluding rx10. since the mortar content in the concrete does not change, it is considered that the compressive strength shows approximately the same value. whether or not the improvement in the air permeability coefficient has been observed in that is whether the part which becomes the air pass has been reformed. in the case of rxs, the strength of the aggregate itself was considered to have increased due to the high splitting tensile strength, but it was considered that the concrete was not modified as rx3 and rx10. 4.4.3. air permeability coefficient in horizontal and vertical direction in order to see the effect of the aggregate interface, the horizontal axis shows the air permeability coefficient in the horizontal direction and the vertical axis shows the air permeability coefficient in the vertical direction in figure 8. in r and rc, it was found that the air permeability in the vertical direction was high, and it is considered that acx bleeding water path was formed, and the air became easier to pass because r and rc had high bleeding amount. on the other hand, the values in the horizontal and vertical directions were almost the same for all rx3, rx10, rxw, and rxs using acx. it is considered that the gap at the aggregate interface was improved by acx. in addition, rc was a modification of the mortar attached to recycled aggregate, so the aggregate interface was not modified. 4.5. relationship with porosity 4.5.1. the relationship between porosity and strength figure 9 shows the relationship between porosity and strength. rx10 improved both strengths over rl, however the pore was almost same. on the other hand, the strength of rxs did not increase and the porosity decreased. for this reason, it is considered that rxs with acx closer to the aggregate also contributed to the modification of the aggregate. rc contributed to splitting tensile strength rather than compressive strength. 4.5.2. the relationship between porosity and air permeability coefficient figure 10 shows the relationship between porosity and air permeability. the rc and rxw obtained by modifying the aggregate have almost the same air permeability as rc, however the porosity decreased. on the other hand, those with the addition of accelerators to concrete, such as rx3 and rx10, improved the air permeability, however the porosity was not improved. therefore, it was found that although the modified mortar portion of the recycled aggregate had a significant effect on the porosity, even if the aggregate was modified, air permeability was not improved. 537 a. shibuya, n. matsuda, t. iyoda acta polytechnica ctu proceedings compressive strength splitting strength dry shrinkage air permeability porosity vertical horizontal aggregate ! ! • ! ! • transition zone • • ! ! • ! mortar • ! • ! ! ! table 3. list of contributions by modification. 4.6. consideration based on the above results, it was considered possible to summarize the thoughts and sort out the degree of influence that each part of concrete has on physical properties. table 3 shows the parts that affect each characteristic. in this study, it was clear that the properties affected by the modified part changed. the properties affected by the respective modifications of the aggregate, the transition zone, and the mortar portion are represented by "•". the parts that do not have an impact are marked as "!", and those that are difficult to judge are marked as "!". 5. conclusion • the quality of recycled concrete improved most when calcium silicate hydrate type accelerator was added to water at 10%. • air permeability can be improved by pre-wetting or spraying with calcium silicate hydrate type accelerator. however, it increases in compressive strength cannot be expected. • forced carbonation of recycled aggregate increases the strength of the aggregate itself and increases the split tensile strength. however, it does not improve the air permeability or compressive strength because it does not modify the aggregate interface in the recycled concrete. • in the future, we would like to study methods for promoting the use of recycled coarse aggregate and study of recycled fine aggregate. in addition, we aim to sort out which part has the greatest effect on the physical properties of concrete. references [1] fukazawa y and iyoda t 2019 clarification of pore modification mechanism in concrete by c-s-h type accelerator 3rd acf symposium assessment and intervention of existing structures. 538 https://doi.org/10.14311/app.2022.33.0564 acta polytechnica ctu proceedings 33:564–570, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague unique and innovative technology for sustainable and efficient structural concrete made of 100% recycled aggregate from cdw bohuslav slánskýa, ∗, pavel zelinkab, jan čermákc a skanska a.s., head of technical research and innovation, křižíkova 682/34a, 186 00 praha 8 karlín, czech republic b skanska transbeton s.r.o., chief technologist, toužimská 664, 190 00 praha letňany, czech republic c eerc-tech, head of r&d, vyšehradská 1349/2, 128 00 praha 2 nové město, czech republic ∗ corresponding author: bohuslav.slansky@skanska.cz abstract. for supporting an environmental and social approach to the construction industry, the authors are bringing a new unique technology for maintaining the production of sustainable and efficient structural concrete made of 100% of recycled aggregate from construction demolition waste (cdw) as a replacement of natural aggregate. a wide range of laboratory testing of concrete with a grade up to c 30/37 using brick and concrete recycled aggregate as well as pilot projects will be presented with very successful results. this has been done by taking a different approach to the concrete mix design and production using new modern materials as well as new methods. then full-scale production testing proved the new concept with only a minor adjustment in concrete plant technology. this new solution has very positive environmental impact and cost optimization shows its great potential for use as a structural concrete supporting a circular economy, saving natural sources, minimizing dumping and landfilling of cdw as well as reducing the co2 footprint. this innovative technology makes the reuse of cdw possible in concrete with higher added value in the construction production chain (upcycling) compared to conservative cdw downcycling. this unique concrete production together with proper cdw management brings with it a quite cost-effective contribution to the building industry and market. keywords: concrete, demolition waste, recycled aggregate. 1. introduction the substitution of natural aggregates in concrete by recycled ones brings both environmental and economic benefits. in terms of environmental protection, it involves the utilization of rarely used, yet recyclable raw materials, which are largely deposited in landfills or dumps, while maintaining these landfilling capacities for materials that cannot be further used. at the same time, it is saving natural aggregate resources for applications where irreplaceable. the reduction of the traffic exposure is another benefit of this method, as the resources of cdw coming from the demolition of old buildings are usually closer to concrete mixing plants than quarries and sand borrow pits usually located far outside of urban areas. in business terms, this method reduces material costs in tens of percent by replacing expensive natural aggregates by what is basically considered a cheap waste. the results of this paper were compared with other research results, where is used 100 % of recycled aggregates. in [1] are compared influences of bricks recycled aggregate on properties of concrete, there is one relevant mixture rmac d with 100 % of recycled aggregates. in mixture rmac d is used 70 kg more cem i 42.5 r than in mixture in table 5. the most significant difference is in used ra, in [1] were used three fractions of ra 0/4, 4/8, 8/16 mm, in this paper is used only mono fraction 0/22 mm, which saves economical and ecological costs during recycling of cdw. the other difference is in parameters of concrete with ra, water absorption is 10 % lower with this technical solution. 2. principles of the technical solution partial substitution of the natural aggregate in concrete by recycled one has been a standard procedure. however, our research and development focused on the production of concrete made solely from the recycled aggregate (ra) including fine fractions of up to 4 mm (100% replacement of natural aggregate including sand). several years of laboratory testing have been finalized with success and new technology has been developed that was patented both in the czech republic as well as internationally. the patented solution used for the production of concrete from cdw has been launched in the czech republic under the registered trademark of rebetongő. the technical solution starts with the proper preparation of ra with demanded granulometry prescribed by technical specifications. the main technical know-how then continues in the production 564 https://doi.org/10.14311/app.2022.33.0564 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 cdw structural concrete process of concrete mixing using "nanofiller", the particles of which reinforce the structure of aggregates itself. the final concrete mixture, in general, achieves similar performance properties as conventional concrete made of natural aggregates, with a strength class of up to 30 mpa including waterproof and frost resistant properties. the use of nanofiller, as well as the reactivation of the binding components contained in recycled aggregate, allows to use the same or even lower cement dose as with quality natural aggregate. due to the softer nature of recycled aggregate, the modulus of elasticity is lower which has to be taken into account in structural design. the used additive ensures the durability of concrete products while allowing for their full recyclability thus completely eliminating future environmental burdens and meeting the strictest views of a circular economy in the construction industry [2, 3]. 3. the development process from the laboratory to industrial production in the concrete plant the performance parameters of concrete made of 100% recycled aggregate are the outcome of the long-term development and laboratory testing in the czech republic with final results in table 1. that was followed by the joint development with the construction company for further industrial optimization and upscale. the first laboratory tests in their laboratories made with concrete mix design in 3 basic modifications using 100% concrete recycled aggregate (rc-c), 100% brick recycled aggregate (rc-b), or mixed recycled aggregate (rc-m) showed surprisingly good results see table 2. this confirmed the high potential of ra concrete for commercial use on construction and development projects. therefore, operational tests have been started immediately at the pilot concrete mixing plant, again with very good results (see table 3). the joint development brought among others the further optimization of production processes, the verification of feasibility in the conditions of industrial production at the concrete plant, and the verification of technological as well as rheological properties of concrete from ra during the production, transport, and casting. based on this good experience the production of concrete made of 100% ra has been certified and since september 2019 branded the certified production rebetong has started. it offers ready-mix concrete in strength classes up to c 25/30, exposure class up to xc4, xf1, and prefabricated wall blocks with the reblok trademark. it focuses on the production of the concrete made of 100% ra mono-fraction 0-22 mm from brick cdw, which makes leftovers in the market and the use of which has a greater economic as well as environmental effect. 4. results of the testing within the framework of research and development, laboratory tests have been performed both on fresh and hardened concrete. as a standard, consistency, air content in fresh concrete, compressive strength of concrete, bulk density of fresh and hardened concrete, have all been performed. on selected samples, concrete bending tensile strength and transverse tensile strength, frost resistance, concrete resistance to de-icing solutions, static modulus of elasticity, water absorption, volume changes of concrete, depth of pressure water infiltration and alkali-silica reaction tests have all been tested. the results of laboratory and production tests show good performance properties of concrete with 100% recycled aggregate made of concrete and bricks. table 1 shows final laboratory result during development phase and table 2 indicate the results of laboratory tests made prior to the industrial production. after laboratory testing, industrial production tests were carried out on the bhs dkxs mixing plant with the mixer volume of 1.67 m3. recipes for concrete strength classes c20/25 and c25/30 with 100% recycled brick aggregate were tested during the first production tests. the test results are indicated in table 3. other concrete parameters tested during certification are shown in table 4, the example of the most commonly used concrete recipes is shown in table 5. one of the success factors is the production of the recycled aggregate from cdw. it took some time and several steps to set up optimize production process, select most suitable equipment and proper set up of each machine in the production chain. development of the concrete made of 100% ra started with 2 fractions 0-8 and 8-16 mm. after some steps of optimization of the ra production the mono-fraction 0-22 mm with similar granulometry was managed. this was an important step for efficient production at concrete plants as they usually lacking of storage space for more fractions of aggregate. also a source of the cdw is important and the best possible way is to manage the whole process starting with selective demolition and create a large storage of the recycled aggregate with similar properties. the final recycled aggregate is tested according en12620 + a1 and furthermore a prescribed limited gradation curve for granulometry is specified see figure 1. storage of the recycled aggregate is similar to natural one. the sensitivity test for the moisture in aggregates between 0 % to full saturation have been done with no significant differences. during the operational tests, the workability of concrete made of recycled aggregate, its behavior, transportability and possibilities of placement have been verified. the behavior of recycled fresh concrete has been found to be similar to conventional natural aggregate concrete. 565 b. slánský, p. zelinka, j. čermák acta polytechnica ctu proceedings sample no. type of rca consistency strength class compressive strength [mpa] 7 days 14 days 28 days 1 rc-c s4 c25/30 18.8 28.4 35.0 2 rc-b s4 c25/30 19.8 27.8 35.3 3 rc-m s4 c25/30 22.5 31.0 38.0 bulk density [kg/m3] tensile bending strength [mpa] tensile splitting strength [mpa] water absorption [%] static modulus of elasticity [gpa] depth of penetration of water [mm] 1 2150 4.7 2.8 3.5 18.6 22.0 2 1920 3.7 2.4 3.2 12.9 16.0 3 2100 4.4 2.5 4.4 14.8 18.0 table 1. results of final laboratory tests during initial development. strength class type of rca consistency sample no. compressive strength on cubes [mpa] bulk density [kg/m3] coefficient of frost resistance 2 d 7 d 28 d 92 d c25/30 rc-c s3 1 11.5 27.2 41.1 45.0 2115 − c25/30 rc-b s4 2 12.2 29.7 44.2 48.5 1913 0.98 c25/30 rc-c s4 3 10.6 25.9 40.0 42.0 2132 1.00 c25/30 rc-b s4 4 13.5 29.3 44.7 48.8 1928 − c25/30 rc-m s4 5 14.4 32.7 46.5 51.7 2043 0.99 c25/30 rc-m s4 6 13.3 30.4 44.4 49.8 2027 − table 2. results of laboratory tests prior start of industrial production. strength class type of rca consistency number of samples compressive strength [mpa] bulk density [kg/m3] depth of water penetration [mm] 7 d 28 d c 20/25 rc-b s4 20 22.0 32.4 1990 19 (6 samples) c 20/25 na s3 11 20.5 28.1 2260 − c 25/30 rc-b s4 6 25.6 36.9 1990 18 (4 samples) table 3. average results of operating tests during industrial production. age of samples [day] compressive strength [mpa] flexural tensile strength [mpa] bulk density [kg/m3] tensile splitting strength [mpa] frost resistance (100 cycles) [−] water absorption [%] static modulus of elasticity [gpa] depth of water penetration [mm] shrinkage [%] c 25/30 with 100 % rc-b 7 24.1 4.3 2020 2.7 0.92 4.1 17.6 16 0.428 14 31.8 0.697 28 37.8 0.977 56 41.0 1.188 c 25/30 with 100 % rc-c 7 19.7 3.9 2110 2.4 0.99 3.9 20.9 21 0.558 14 22.8 0.854 28 35.1 1.119 56 37.5 1.307 table 4. results of concrete parameters during certification. 566 vol. 33/2022 cdw structural concrete component quantity [kg/m3] 100% rc-b 100% na cem i 42,5 r 250 270 ground limestone / or fly ash 50 50 nanofiller (microsilica) 25 − natural sand 0-4 mm − 870 natural crushed aggregate 4-8 mm − 250 natural crushed aggregate 8-16 mm − 700 brick recycled aggregates 0-22 mm 1500 − admixture (plasticizer) 5,5 2,0 water 210 165 table 5. comparison of concrete mix design for c 20/25 with ra and na. figure 1. stress diagram of a strengthened beam. source: adapted from silveira [4]. strength class type of ra time [min] air content [%] slump test [mm] temperature of concrete [◦c] c25/30 rc-b 0 4.4 210 23.3 30 4.2 210 23 60 4.0 160 22.5 table 6. results of workability testing in time. �� � )ljxuh����3lfwxuhv�iurp�erqg�vwuhqjwk�whvw��figure 2. pictures from bond strength test. 567 b. slánský, p. zelinka, j. čermák acta polytechnica ctu proceedings )ljxuh���� � � )ljxuh����6kulqndjh�ghyhorsphqw�whvw��ohiw�±������eulfn�5$��uljkw�±������1$�� figure 3. shrinkage development test: left 100% brick ra, right 100% na. ��� � figure 4. concrete placement during the first semi-operational tests and concrete blocks. figure 5. the first ready mix concrete placement. �� � figure 6. residential project devil’s hill in prague and first concrete wall. 568 vol. 33/2022 cdw structural concrete due to the higher fluctuation of the absorption of ra, especially the one consisting of bricks, it is necessary to pay attention to the workability time of the fresh concrete and to use special additives to ensure good workability for at least 60 minutes. within the testing, consistency using the slump test was determined over a time horizon of sixty minutes on brick recycled concrete to verify the workability and transportability of recycled concrete. the results are displayed in table 6. after satisfactory standard test results some additional over-standard tests, according to [5], were carried out for concrete made of 100% brick recycled aggregate. for example reinforcement bars bond test according to [4] with final bond strength of 5,2 mpa compare to reference result of 5,4 mpa for a concrete made fully of natural aggregate (na). pictures from the test see in figure 2. other additional test was the shrinkage development test according to [6]. resulting graphs are in the figure 3. 5. pilot projects, application examples the concrete made from 100% ra has performance properties similar to the conventional concrete made from natural aggregates. therefore, it can be used for common plain and reinforced concrete structures. when designing horizontal load-bearing structures, consider the lower modulus of elasticity (see chap. 4.). customers have become soon interested in the prefabricated wall blocks, of which almost 2000 pcs have already been produced see figure 4. as a presentation of the pilot production, these precast units were used for the construction of the parkour playground at "cukrkandl" in prague. the first pilot project of a ready-mix concrete made of 100% rc-b was executed for a base concrete for a commercial multi storey project in olomouc in a volume of 350 m3 see figure 5. the next complex pilot project was a residential project called devil’s hill in prague, where 2000 m3 of 100% rc-b were placed for base slabs and reinforced concrete walls see figure 6. 6. technical legislation currently, not only within the czech republic but throughout the eu, there are significant barriers in technical regulations that significantly reduce the use of recycled aggregates in concrete. although properties of ra are specified in european standard en12602 + a1, the european standard for concrete en206 + a1 recommends the use of coarse recycled aggregate in concrete only up to a maximum of 50% for types a and b, and only for the x0 exposure class, and even less for other exposure classes. to support wider use of ra in concrete, it is necessary to change this conservative normative approach and focus on proving the final properties of the concrete without limiting the ra content. in the czech republic, this process, thanks to significant support from the ministry of the environment, and ministry of industry and trade, started in september 2019 and its aim is to create a revised national standard for concrete with up to 100% of ra (čsn p 73 2404). at the time being the only possible way to apply 100% recycled aggregate in concrete is the procedure according to the national act no. 163 i.e. certification of the system of production control of concrete from ra and introducing the product to the market on a basis of building technical approval (sto). unfortunately, this procedure cannot be applied to projects where the contractor is contractually bound by applicable national and european standards. at the moment, this applies for the majority of public procurement projects. however the situation is slowly but surely changing as public investors are increasingly becoming aware of the importance of having a sustainable construction in public procurement and a circular economy. recently they are coming up with requirements in their projects for construction products made from secondary, recycled raw materials. 7. conclusion the parameters determined during laboratory and operational testing show a great potential of concrete made from 100% recycled aggregate, since its performance properties are comparable with conventional concrete in most indicators. the patented solution is in full compliance with the principles of a circular economy, where all inert cdw can be converted back into the building material with excellent properties, which is in addition fully recyclable using the same technological process. considering the ever-decreasing stocks of natural aggregates, the use of alternative resources is economically and environmentally beneficial. the use of a patented technological process also results in a reduction in the carbon footprint by more than 10%, as less cement is used for the production of concrete and the otherwise rarely used cdw are used as a replacement of natural aggregates. references [1] t. pavlu, k. fortova, j. divis, et al. the utilization of recycled masonry aggregate and recycled eps for concrete blocks for mortarless masonry. materials 12(12), 2019. https://doi.org/10.3390/ma12121923. [2] j. čermák, j. fiala, f. polak. comparison of utility properties of concrete with different types of recycled aggregates. brno: brno university of technology, faculty of mechanical engineering, 1996. [3] b. slánský, p. zelinka, j. čermák. beton z recyklovaného kameniva. in: betonářské dny: konference s mezinárodní účastí. pardubice: česká betonářská společnost, 1993. 569 https://doi.org/10.3390/ma12121923 b. slánský, p. zelinka, j. čermák acta polytechnica ctu proceedings [4] čsn en 73 2404. concrete: specification, performance, production and conformity additional information. prague: czech normalization institut, 2016. [5] čsn 73 1328. determination of shear adhesion of steel to concrete. prague: czech normalization institut, 1971. [6] önorm b 3329. grout: requirements and test methods. austrian standards, 2009. 570 https://doi.org/10.14311/app.2022.33.0052 acta polytechnica ctu proceedings 33:52–57, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague application of high performance concrete for reconstruction iva broukalová∗, alena kohoutková, anna horáková, hana schreiberová czech technical university in prague, department of concrete and masonry structures, faculty of civil engineering, thákurova 2077/7, 166 29 praha 6 dejvice, czech republic ∗ corresponding author: iva.broukalova@fsv.cvut.cz abstract. one of the sustainability aspects of the construction is a sufficient durability. in objectively justified cases, the durability is enhanced by reconstruction. to ensure economic design of the reconstruction, non-liner approaches are often applied. this paper deals with aănonlinear numerical simulation of disassembled reinforced concrete (rc) panels strengthened by an ultra-high performance concrete (uhpc) layer. based on the prior experimental programme, simulations of four-point bending tests of the original and strengthened panels are created using software atena science. calibration of the material parameters is based on the destructive and non-destructive investigations performed in the experimental programme. the comparison of the experimental and numerical model loading curves indicates that additional mechanical testing will be needed in order to achieve an accurate numerical simulation. although the bending test of the rc panel and the crack mouth opening displacement (cmod) test performed on the applied uhpc beam were calibrated, the final model of the uhpc strengthened panel does not correspond completely with the experimental measurements. in this paper the possible reasons for this result are discussed. keywords: analysis, high-performance concrete, reconstruction. 1. state of the art in recent decades, environmental aspects have been attracting more and more attention in the field of building structures. one of the possible approaches to an environmentally friendlier structure is prolongation of its service life. by improving the no-longer complying parts of structures or further utilization of removed building components, the negative impacts of construction can be reduced greatly. however, an increased demand for the load-bearing capacity or a different loading type of the recycled component very often results in necessary strengthening of the structure. in the case of concrete structures, an application of a new additional concrete layer, usually from a type of high-performance fibre concrete, has been proposed and investigated as a potential strengthening method [1–13]. in the previous studies, the layer from highperformance (hpc) or ultra-high performance (uhpc) concrete has been generally applied in four strengthening configurations in thickness from 20 to 100ămm (most frequently 30 mm). a strengthening layer located in the compression zone was most frequently used in the case of panels [2, 5, 8, 12]. the main advantage of this configuration lies in its easier performing as the strengthening layer is applied on the upper surface, thus no complex formwork is needed. concrete beams have been, in the conducted experiments, strengthened in the compressive zone [1, 4, 10, 11], the tension zone [3, 7, 11], two-sided (left and right surfaces) [7, 13], and three-sided (aăcombination of all of the above) [3, 7, 9, 11, 13]. in order to determine the most effective strengthening configuration, a number of studies compared the impact of the layer location on the load-bearing capacity of the element [7, 11]. the results indicated that the three-sided configuration reaches the highest values (an increase of up to 89% in the loadbearing capacity) following by the two-sided (around 47%) and the tensile configuration with as low as 16%. not surprisingly, the conducted experiments also showed that the magnitude of the tensile strength of the strengthening layer has a decisive influence especially in the three-sided configuration, whereas it has only a little effect in the case of compressive zone [7]. another important aspect of the hpc or uhpc strengthening method is the adhesion of the new layer with the original concrete component. in [7, 13] which have dealt with this topic, a transverse tensile strength test was performed on cylindrical or cubic specimens which were prepared from both the original basic concrete and uhpc. the results of these experiments were also further used in numerical finite element (fe) modelling in order to describe the behaviour of the contact layer. to further improve the adhesion of the layers, researchers have been also investigating different surface treatments, such as certain adhesion agents 52 https://doi.org/10.14311/app.2022.33.0052 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 application of high performance concrete for reconstruction figure 1. a typical geometry of the non-strengthened panels based on destructive investigations. figure 2. typical geometry of the strengthened panels after the uhpc layer application [14]. (e.g. epoxy resin) or roughening (e.g. sandblasting, hydrodemolition, or wire-brushing). a study [7] showed, that specimens with an epoxy resin surface treatment reach higher values in the transverse tensile strength test but lower values in the shear strength test compared to specimens with a surface treated with sandblasting. importantly, the further bending test showed that the sandblasted surface leads to higher load-bearing capacity compared to the epoxy treatment. the finding indicates that the strength of the new-to-old concrete bonding depends primarily on its shear strength. this paper focuses on numerical modelling of the uhpc strengthening method. as high-performance concrete is rather economically and ecologically demanding material, careful optimization of its placement and characteristics is crucial for the sake of maintaining the positive effects of the application of the method. in order to achieve an effectively strengthened concrete structure by an uhpc layer, it is necessary to describe characteristics of the materials and their interaction as accurately as possible. for this reason, a large number of studies have been conducting numerical non-linear analysis alongside mechanical testing [7, 11–13]. by using an appropriate finite element software (e.g., atena, abaqus) and material models, the studies have been generally able to successfully describe the structure behaviour with only little (no more than 13.4%) deviations from the mechanical test results. however, several issues in the topic of fe modelling needs further clarification as they might influence the model accuracy. firstly, the compressive and tensile stress-strain relationship for the uhpc is a rather complex problem, as it is significantly different from the normal strength concrete. further, the characteristics of the bonding between the original substrate and new uhpc layer may distort the results, although the majority of previous studies considered perfect bonding. in our paper, we focus on fe numerical analysis of four-point bending tests performed on prefabricated reinforced concrete panels strengthened by a uhpc layer. in this study, several calibration procedures of material parameters are presented, and the outcomes are compared with experimental measurements. further, the specifics of designing recycled structures are discussed and further development is outlined. 2. experimental program 2.1. specimens bending tests were performed on several prefabricated reinforced concrete (rc) panels, which have been located in outside conditions for several years. the panels were around 120 mm high, 490 mm wide, 53 i. broukalová, a. kohoutková, a. horáková, h. schreiberová acta polytechnica ctu proceedings and 2750 mm long; however, the dimensions varied slightly in every specimen. destructive tests showed that the panels were reinforced with steel reinforcement two !12 mm bars at the bottom surface and two !12 mm bars at the upper surface. however, the performed destructive investigations revealed that the location of the bars was slightly different in every specimen. further, due to the inadequate reinforcement at the bottom surface, four !10mm b500b bars were glued into formed grooves to increase the loadbearing capacity of the panels. before casting of the uhpc layer, the upper surfaces of the panels were treated by hand hydro-jetting to ensure adequate bonding of the original substrate with the new layer. after the treatment, the uhpc layers were cast in the compressive zone in thickness of 30 mm. this study focuses on two experimental series reference panels with no strengthening layer (figure 1) and panels with 30 mm uhpc layer (figure 2). 2.2. material properties in order to determine the strength class of the original rc panels, a number of destructive and nondestructive tests were performed. according to the measurements, the material was categorized as c30/37. the material characteristics of the prepared ultrahigh strength fibre concrete (uhpc) were investigated by destructive mechanical testing. the compressive test conducted on cubic specimens determined its mean compressive strength as 133 mpa. further, a three-point bending test, with support span 500 mm, was conducted on pre-notched beam specimens (150 × 150 × 700 mm) to obtain curves of the dependence of the loading force on the crack mouth opening displacement (cmod). 2.3. testing methods to determine behaviour of the investigated strengthened and non-strengthened panels, a series of fourpoint bending tests were conducted. the span of the supporting pins was 2500 mm, the span of the loading pins was 800 mm. the loading force and deformation were monitored and recorded during the loading. 3. numerical modelling of the four-point bending tests all of the analysis was conducted using a finite element analysis software atena 5.6.1 science with gid interface. the actual geometry of the test setup was modelled in all cases. in the case of the fourpoint bending tests, the symmetry of the setup was considered, and only halves of the specimens were modelled. in all cases, the loading was introduced into the model as a pre-described deformation through steel distribution plates. 3.1. numerical modelling of the reference non-strengthened panels firstly, the four-point bending test of the nonstrengthened panels was modelled. as the information about its materials was limited, a careful calibration of the numerical model was crucial for further investigation of the uhpc strengthening method. in order to determine whether it would be possible to omit the spatial effect, the rc panel was modelled as both a plane stress (2d) and a three-dimensional stress (3d) structure model type. the geometry was identical in both cases. however, due to the significantly lower computational demands of the 2d model, the mesh size could be set to 15 × 15 mm in the case of the 2d model, whereas the 3d model used 25 × 25 mm. in order to obtain the loading curves, the loading force was monitored using a monitor placed under the force and the deflection was monitored on the lower surface in the middle of the span. 3.1.1. material models as mentioned above, the information about the rc original materials were limited. as a starting point, the material parameters of both concrete and steel bars were generated using the predefined materials in atena gid software. material concrete ec2 which uses the cc3dnonlincementitous2 material model was selected as a representation of concrete. as this material generates the parameters according to the eurocode2, the category c30/37 was set. the steel bars, both the original and the additional, were defined as reinforcement ec2 using ccreinforcement set to the category b500b. 3.1.2. parametric study and calibration of the fe models firstly, analysis of the 2d and 3d model with generated default material parameters, using the material models mentioned above, was run. subsequently, the loading-deflection curves obtained from the analysis were compared in order to evaluate the differences between the 2d and 3d modelling. figure 3. a comparison between the 2d and 3d loading curve a reference panel with default material parameters. as can be seen in figure 3, generally, the curves did not differ significantly. the crack formation and the 54 vol. 33/2022 application of high performance concrete for reconstruction parameter units default settin multiplication factor concrete material model young modulus mpa 32000 0.56 tension strength mpa 2.9 1.00 compressive strength mpa −38 0.79 fracture energy mn/m 0.000073 1.14 tension stiffening − no yes critical compressive displacement m −0.0005 1.10 original steel reinforcement model characteristic yield strength mpa 500 0.84 class − b c table 1. quality and operations assessment schedules [7]. figure 4. the experimental loading curves of the tested non-strengthened panels compared with the calibrated atena simulation. reinforcing bars yield occurred at almost the identical loading force (at approx. 15 kn and 63ăkn, respectively). however, before the yield point, the lower slope of the 2d curve indicated a slightly lower stiffness in the case of 2d analysis, especially after the crack formation. on the other hand, after the yielding, the slope of the 2d curve was higher compared to the 3d curve. furthermore, the failure of the 2d model took place at significantly higher values of the deflection, resulting in a higher maximum load force. on the base of a parametric study (the analysis was run several times, while changing the individual values of the default material models) the numerical models were calibrated (figure 4) by setting the material parameter values to required values, so that the loading curve obtained from the fe analysis would correspond as closely as possible with the experimental measurements. in table 1, the key default parameters together with the necessary multiplying coefficient for the changed values are listed. 3.2. numerical modelling of the panels with a uhpc layer after the calibration of the model of the nonstrengthened panel, analysis of the strengthened specimens could take place. conveniently, unlike the original materials, the newly introduced uhpc could have been subjected to several destructive tests. thus, the information about its compressive strength figure 5. the experimental force-cmod curve compared with the fe analysis results (with default material parameters and after the calibration procedure). and tensile behavior was much more certain. to correctly determine its tensile material parameters, a three-point bending test was modelled in the atena software and calibrated using the conducted experiment. subsequently, the strengthened panel was modelled using the parameters obtained from the numerical representation of the bending test. 3.2.1. the cmod analysis of the applied uhpc in order to define the uhpc material parameters, aăuhpc beam subjected to the three-point bending test was modelled using atena science in 2d. in order to obtain the force-cmod curves, the loading force and horizontal displacements in the notch mouth were monitored. as a representation of the uhpc, cc3dnonlincementitous2user was utilized as it allows the user to define its own fracture-plastic material model, and it is recommended for the fibre reinforced concrete (frc) modelling. the calibration procedure was performed according to the software documentation. the tensile function, which have a crucial importance for the frc materials, were carefully set so that the force-cmod curves corresponded and closely as possible (figure 5). calibrated values of the material parameters can be seen in table 2. 55 i. broukalová, a. kohoutková, a. horáková, h. schreiberová acta polytechnica ctu proceedings parameter units calibrated values parameter units calibrated values young modulus mpa 43000 tension strength mpa 6.5 compressive strength mpa −133 table 2. calibrated values of the key material parameters based on the cmod analysis. 3.2.2. modelling of the four-point bending test of the strengthened panels based on the four-point bending test of the strengthened panels, the strengthened panels were modelled using atena science as a plane stress structure model type. the geometry of the model remained identical, only the 30 mm uhpc layer was added. the connection between the old and new layer was considered perfectly solid, thus no mechanical characteristics were prescribed. the material parameters of the steel bars and original concrete were set according to the calibration of the non-strengthened panels. the uhpc parameters were taken over from the calibrated cmod analysis. results of the analysis run with the parameters based on the prior calibration procedures can be seen in figure 6. figure 6. the experimental loading curve of strengthened panels compared with the atena simulation with material parameters based on the prior calibration procedures. as can be seen, the load-deflection curve from obtained from the fe analysis did not match the experimental measurements completely. similarly, the formation of the tensile cracks occurred around the same loading force and deflection. however, the stiffness of the cross-section after the cracking differed quite significantly, as the curve slope of the fe model is notably steeper. the fe model also showed approx. 12 % higher maximum loading force than the experimental measurements. 4. discussion the finite element analysis of strengthened prefab concrete panels presented in this paper pointed out several crucial factors which influence the design of the uhpc strengthening method. as previously stated in the introduction, the environmental benefits of recycling or renewal of no-longer necessary or unsatisfactory concrete structures are unquestionable. however, as is apparent from the experimental part of the study, the knowledge of the original materials and its geometrical arrangement in the structure is frequently incomplete. for this reason, the numerical analysis inputs are then based on several destructive or non-destructive tests, which provide only limited information about selected specimens. as the parametric study of the non-strengthened panels conducted in our paper showed, the structure behaviour was fundamentally affected by the placement and characteristics of the original steel reinforcement. in our experimental program, the placement, and possibly the diameter, of the bars were slightly different in every specimen, and it was not possible to subject the bars to any mechanical measurements. thus, although the fe model of the fourpoint bending test corresponded well with the experiment after the calibration, it was adjusted only to the chosen specimens. therefore, the model could have been unsatisfactory for the further strengthened panels due to a different geometrical arrangement. more known inputs were available for the modelling of the strengthened panels. the cmod analysis of uhpc provided information about the material tensile properties and the compressive test determined its compressive strength. however, no loaddeflection curve was known from the pressure test, only the maximum loading force. thus, the compressive function (the stress-strain relationship) could not have been determined based on any measurements. the loading curve of the strengthened panel showed, that the calibrated model of the original panel and uhpc with defined tensile function based on the cmod analysis did not match with the experimental measurements. this unsatisfactory result could have been caused by several factors. firstly, as stated above, the model fitted on the selected non-strengthened panels could have been unsuitable for the selected strengthened panels due to the different geometry or materials. secondly, as mentioned in the introduction, the tensile properties of the strengthening layer in the compressive zone have only a limited impact on the load-bearing capacity. thus, the setting of the compressive function based on the load-deflection curve could have been crucial and more experimental measurement is therefore needed. further, no stresses due to shrinkage of the uhpc layer and mechanical properties of the connection between the original substrate and new material were considered. although no visible deterioration of the layers was apparent during the bending test, the effect may not be negligible in the fe analysis and determination of the interface parameters will be necessary in future investigations. 56 vol. 33/2022 application of high performance concrete for reconstruction 5. conclusions application of high-performance cementitious composites for reconstruction of concrete structures is aăstep towards sustainable construction. the reuse of damaged or no-longer satisfactory structural elements leads to saving both material and waste. however, an optimized design of the required strengthening is the key to ensure its advantageousness. when the non-linear behaviour of the materials is taken into account, it allows the design to be less conservative, thus more economic. in this paper, the performed non-linear numerical analysis pointed out several crucial factors which influence the numerical simulation suitability. for a sufficient numerical model, the geometry of the recycled structures must be known as precisely as possible. further, when modelling only a general geometry of the structures, the slight differences in the individual elements might affect the model accuracy greatly. however, most importantly, it is crucial to ensure inputs for the numerical analysis by conducting a sufficient number of mechanical tests of the materials present in the modelled structure. acknowledgements fv20472 application of high performance cementitious composites at rehabilitation of concrete structures, sgs19/149/ohk1/3t/11 durability of concrete structure and assessment of its life cycle. references [1] m. genedy, j. kim, m. r. taha. innovative strengthening of rc beams using cfrpuhpc composite. proceedings of the 7th international conference on frp composites in civil engineering, vancouver, 2014. [2] t. boeters, r. braam, h. kolstein, et al. concrete overlay of movable steel orthotropic bridge decks. steel construction 2(2):104-8, 2009. https://doi.org/10.1002/stco.200910014. [3] f. a. farhat, d. nicolaides, a. kanellopoulos, et al. behavior of rc beams retrofitted with cardifrc after thermal cycling. journal of materials in civil engineering 22(1):21-8, 2010. https: //doi.org/10.1061/(asce)mt.1943-5533.0000008. [4] m. genedy. a new cfrp-uhpc system for strengthening reinforced concrete t-beams, master thesis, 2014. https://digitalrepository.unm.edu/cgi/viewcont ent.cgi?article=1094&context=ce_etds. [5] m. bastien-masse, e. brühwiler. ultra high performance fiber reinforced concrete for strengthening and protecting bridge deck slabs. bridge maintenance, safety, management and life extension, p. 2176-82, 2014. https://doi.org/10.1201/b17063-333. 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[10] experimental investigation on reinforced ultra-high-performance fiber-reinforced concrete composite beams subjected to combined bending and shear. aci structural journal 110(2), 2013. https://doi.org/10.14359/51684405. [11] a. p. lampropoulos, s. a. paschalis, o. t. tsioulou, et al. strengthening of reinforced concrete beams using ultra high performance fibre reinforced concrete (uhpfrc). engineering structures 106:370-84, 2016. https: //doi.org/10.1016/j.engstruct.2015.10.042. [12] j. moon, m. m. reda taha, j. j. kim. flexural strengthening of rc slabs using a hybrid frp-uhpc system including shear connector. advances in materials science and engineering 2017:1-7, 2017. https://doi.org/10.1155/2017/4387545. [13] a. a. bahraq, m. a. al-osta, s. ahmad, et al. experimental and numerical investigation of shear behavior of rc beams strengthened by ultra-high performance concrete. international journal of concrete structures and materials 13(1), 2019. https://doi.org/10.1186/s40069-018-0330-z. [14] a. čítek. experimentální ověření betonových prvků zesílených pomocí uhpc, čvut v praze, fakulta stavební, 2019. 57 https://doi.org/10.1002/stco.200910014 https://doi.org/10.1061/(asce)mt.1943-5533.0000008 https://digitalrepository.unm.edu/cgi/viewcontent.cgi?article=1094&context=ce_etds https://doi.org/10.1201/b17063-333 https://doi.org/10.1016/j.conbuildmat.2017.08.077 https://doi.org/10.1016/j.conbuildmat.2016.12.094 https://doi.org/10.1088/1757-899x/271/1/012064 https://doi.org/10.14359/51684405 https://doi.org/10.1016/j.engstruct.2015.10.042 https://doi.org/10.1155/2017/4387545 https://doi.org/10.1186/s40069-018-0330-z https://doi.org/10.14311/app.2022.33.0604 acta polytechnica ctu proceedings 33:604–609, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the status of sustainable concrete codes in the united states andrew warren taylor kpff consulting engineers, 1601 fifth avenue, suite 1600, seattle, wa 98117, usa correspondence: andrew.taylor@kpff.com abstract. as the focus on embodied carbon within the built environment has grown in the u.s., local jurisdictional codes have increasingly proposed and implemented code language that requires the use of concrete mixtures that have reduced environmental impacts. this paper provides details about the various ways that localized sustainable concrete provisions have been proposed and implemented in the u.s., including limits on cement content and global warming potential (gwp). these laws and code revisions vary in their measurement of low-carbon concrete and their exemptions and incentives. while the current version of the aci building code, aci 318-19, does allow for the consideration of sustainability in concrete design, specifics about how the relative sustainability of concrete mixtures and systems are to be measured and compared are not included. this paper addresses how the changing landscape of local jurisdictional codes is driving the development of future aci 318 provisions to play a part in providing more consistency in the application of sustainable design practices in the united states. keywords: codes, concrete, sustainability. 1. introduction in the last decade engineers, architects, local and national government agencies, and national building code committees in the united states have started to evaluate and take measures to mitigate the environmental impacts of building construction and operations. as these efforts mature, new programs to encourage more sustainable design and operation of buildings, and new regulations that require the reduction of carbon emissions, have been established. regarding concrete structures, it is most common for new programs and regulations to focus on the carbon emissions associated with the initial construction of a building: atmospheric carbon released by the manufacturing of portland cement, by the mining and transportation of constituent materials, and by the construction process itself. less often, consideration is given to the carbon emissions associated with the complete life-cycle of the building: the construction period; the operating life of the building; and the demolition and disposal or recycling of the building materials. when viewed in this broader life-cycle sense, the carbon emissions associated with portland cement manufacturing make up a smaller proportion of the total atmospheric carbon impact than when only the initial construction emissions are considered. this paper describes the most recent programs and regulations implemented in the united states that promote the reduction of carbon emissions associated with concrete buildings. programs and regulations that consider only the atmospheric carbon emissions related to initial construction, as well as one program that consider the impacts of the entire life cycle of the building, are summarized 2. metric for global warming potential the de facto international standard for assessing the global warming potential (gwp) of construction materials is the environmental product declaration (epd). an epd is prepared in accordance with iso 14025 [1], and each epd provides a quantitative measure of gwp in terms of kilograms of co2 per cubic meter of material (kg co2e/m3). an epd is required for each variation of a construction material; thus a unique epd is necessary for each concrete mixture that is incorporated into a building project for which a carbon emissions assessment is performed. the existence of epds, and gwp ratings provides a consistent basis for evaluating the overall gwp of a project. epds and gwp ratings also enable government agencies and code-writing authorities to establish objective frameworks for quantifying the greenhouse gas (ghg) emissions associated with a particular building project, group of projects, type of construction, or regional or national construction activity. the gwp of a particular concrete mixture depends on many factors, including portland cement content, supplementary cementitious materials content, coarse and fine aggregate sources, and the density of aggregates (normal weight vs. light weight aggregates). in general, the gwp of most concrete mixtures commonly produced in the united states varies between about 200 and 700 kg co2e/m3 604 https://doi.org/10.14311/app.2022.33.0604 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 sustainable concrete codes in the usa 3. organizations promoting reduced gwp of concrete in the united states several organizations have emerged as leaders of the movement to reduce the environmental impacts of concrete. generally, these are not-for-profit organizations, professional societies, or industry consortia. four organizations that support a reduction in gwp of concrete are described below. 3.1. carbon leadership forum the carbon leadership forum (clf) is a not-forprofit forum operated within the college of built environments at the university of washington, in seattle. the clf was founded in 2009 with the objective of developing and promoting quantitative and standardized measures of embodied carbon in construction materials. a primary focus of the clf has been embodied carbon of construction materials, including the promotion of standardized and verified epds, which will allow valid comparisons between various materials and between various suppliers of a given material. a major accomplishment of the clf has been to foster development of computer software that provides a standardized method for calculating the embodied carbon in construction materials. this software was released on november 19, 2019 and is called the embodied carbon in construction calculator, or ec3. the ec3 software is available without charge. it relies on a database of over 19,000 epds from the united states and canada. the vast majority of the epds in the database are for concrete mixtures; epds for other materials as well as additional epds for concrete are continuously added as they become available [2]. 3.2. structural engineers 2050 challenge the structural engineers 2050 challenge (se 2050) was conceived in 2016 as a way to encourage the structural engineering profession to actively participate in a movement towards buildings with greatly reduced, or zero, net embodied carbon. since structural engineers are responsible for the selection of structural materials, structural engineers are in the best position to implement structural materials with the lowest possible embodied carbon, and to report embodied carbon of individual projects to a centralized database. it is envisioned that such a database will be used to track annual total embodied carbon for buildings in the united states. membership in the challenge is voluntary. as of this writing, signatories to the se 2050 challenge include 19 structural engineering consulting companies, 62 individual structural engineers, and 32 other individuals or companies affiliated with the building design and construction industry. 3.3. architecture 2030 challenge the non-profit group architecture 2030 was founded in 2002 to address the need for reducing operating carbon emissions from the built environment. in 2006 architecture 2030 launched the 2030 challenge, which defines quantitative goals for reducing the greenhouse gas emissions associated with the ongoing operations of buildings. emphasis is placed on reducing the consumption of energy generated by fossil fuels, with the goal of reducing operational greenhouse gas emissions of all new or renovated buildings to zero by 2030. "these targets may be met by implementing innovative sustainable design strategies, generating on-site renewable energy, and/or purchasing (20% maximum) off-site renewable energy." [3]. to support the architecture 2030 challenge, the american institute of architects (aia) has created the "2030 commitment." this is a program under which signatory design firms voluntarily report project data, in a standardized format, which quantifies the operational greenhouse gas emissions over the life of the building. reported data is collected in a central database called the design data exchange (ddx). the ddx can be searched using multiple project attributes (like building type, size, and environmental conditions) to help designers create buildings with reduced operational carbon emissions. currently over 600 design firms have signed on to the 2030 commitment, and each year project reports are added to the ddx for hundreds of millions of square meters floor area of new and renovated buildings. 3.4. the american concrete institute in recent years the american concrete institute (aci) has sought to address the challenge of reducing greenhouse gas emissions through a range of initiatives and publications. these include more than 40 standards and guides containing information on sustainability in concrete construction; over 60 journal articles related to sustainability; 200 practiceoriented papers and articles on sustainable design and construction methods; more than 100 online learning tools; and sponsorship or co-sponsorship of over 25 events that have supported the advancement of sustainability in concrete design and construction. in 2008, the american concrete institute (aci) initiated a focus on mitigating the environmental impacts of concrete with the formation of technical committee aci 130 "sustainability of concrete." the committee currently has over 100 voting and associate members, and maintains eight subcommittees: • materials • production/transportation/construction • structures in service • rating systems/sustainability tools • design/specifications/codes/regulations • education 605 andrew warren taylor acta polytechnica ctu proceedings figure 1. local government and state initiatives related to sustainability of construction materials. • climate change impacts on the sustainability of concrete • liaison subcommittee in 2019 committee 130 published report on the role of materials in sustainable concrete construction [4] which provides comprehensive background on sustainability topics, and describes available practices for design, production, and construction that reduce the environmental impacts of concrete. aci formed innovation task group 10 (itg-10), which in 2019 published two important reports that provide guidance on the characteristics and implementation of alternative cements (i.e. alternatives to portland cement): report on alternative cements [5], and practitioner’s guide for alternative cements [6]. these reports present practical approaches to incorporating alternative cements in concrete mixtures, as a means of reducing greenhouse gas emissions caused by production of portland cement. it should be noted that in the 2019 edition of the aci 318 building code [7], section 26.4.1.1.1 was added with the intention of providing increased flexibility to engineers and contractors for incorporating alternative cements in concrete construction: "alternative cements shall be permitted if approved by the licensed design professional and the building official. approval shall be based upon test data documenting that the proposed concrete mixture made with the alternative cement meets the performance requirements for the application including structural, fire, and durability." in order to further develop sustainability provisions in the aci 318 code, in 2019 a new technical subcommittee of the building code committee was established: subcommittee n, "sustainability." this new subcommittee is currently working on revisions to the 2025 edition of the aci 318 code, including an appendix with expanded provisions regarding sustainability. 4. local and state initiatives to quantify and reduce gwp of concrete in the united states within the last two years several local and state government initiatives have been either proposed or enacted to reduce the environmental impacts of concrete construction. these initiatives are identified in figure 1. in addition, the author is aware of preliminary discussions within other jurisdictions that may lead to further concrete sustainability initiatives, but it is still too early to report on the details of those efforts. for example, the marin county ordinance (figure 1, and section 4.3) is likely to serve as a model ordinance for adoption by other san francisco bay area governments. 4.1. the buy clean washington act on january 8, 2018 a proposal was introduced to the washington state house of representatives, house bill 2412, titled the "buy clean washington act." [7] this proposed legislation recognizes that "climate change will have devastating global impacts," and that "all scientific evidence points to the need for washington and the world to reduce greenhouse gas emissions to avert the worst effects of climate change." to reduce greenhouse gas emissions, the legislation proposes "incorporating emissions information from throughout the supply chain and product life cycle into procurement decisions, and using that information to help direct expenditure." in other words, if adopted, the legislation would make global warming potential a key factor in the selection of construction materials and suppliers in washington state. this act would apply to new construction with a floor area greater than 500 square meters and a building renovations greater than 500 square meters where 606 vol. 33/2022 sustainable concrete codes in the usa the cost of the renovation is greater than 50 percent of the value of the building. the list of construction materials that this act would apply to includes • carbon steel concrete reinforcing bars • flat glass • mineral wool board insulation • structural steel • cement • structural timber • solar panels • refrigerants in new equipment • aluminum • gypsum • concrete the proposed legislation was not adopted during the 2018 washington state legislative session, but it is likely the proposal will be re-introduced in 2020. 4.2. city of portland ordinance the city of portland, in the state of oregon, enacted an ordinance in january 2020 that is intended to be the starting point of regulations to limit the greenhouse gas emissions of concrete purchased by the city. at this time the ordinance requires reporting of a product-specific type iii environmental product declaration that is verified by a third party and is less than 5 years old for every concrete mixture supplied for city-funded projects. the ordinance does not, however, currently set limits on embodied carbon or global warming potential of concrete. such limits will be published by the city by april 1, 2021, when the city will issue limits on the maximum global warming potential of any concrete mixtures on city-funded projects. the ordinance does not apply to small projects that require less than 38 cubic meters of concrete. 4.3. marin county ordinance one of the most significant recent developments in regulatory rules regarding the global warming potential of cement and concrete is the enactment of an ordinance in marin county, california. this ordinance, titled "low carbon concrete requirements," forms section 19.07 of the marin county code [9]. it was adopted in november, 2019. in establishing the rationale for implementing this ordinance, the county board of supervisors stated, ". . . marin county experiences climatic seasonal reduction in vegetative moisture content, combined with our heavily populated steep terrain, which presents increased wildfire risk to our residents from carbon-induced global warming; and is also bordered by sea water on three sides and subject to direct adverse local impact from sea-level rise as the result of construction related contributions to climate change, including significant carbon emissions from cement production." to accomplish the goal of reducing greenhouse gas emissions by controlling embodied carbon in concrete, marin county has taken the approach of providing two alternatives: all concrete used in marin county must either meet maximum limits on cement content, or it must meet maximum limits on embodied carbon. these limits are dependent on the minimum specified compressive concrete strength (f ′c). table 1 below is adapted from table 19.07.050 of the marin county ordinance [9]. an exception to the limits in table 1 is permitted for concrete that must attain a high early strength, such as concrete used in precast or pre-stressed construction, concrete used in above-ground beams and slabs, and pneumatically placed concrete (shotcrete). in those cases the values in table 1 may be exceeded by 30 percent. another exception is allowed to the limits in table 1 if a pre-approved cement is used that has an environmental product declaration with embodied carbon less than 1,040 kg co2e per metric ton. the marin county regulations permit compliance with the table 1 limits on the basis of either a) the concrete mixture design, or b) based on the total volume of concrete used in the project. the latter method provides a way to use a variety of concrete mixtures on a project, some of which may not meet the table 1 limits, provided the total cement usage or the total embodied carbon of concrete used on the project does not exceed the hypothetical total cement usage or the total embodied carbon of concrete calculated based on the actual volume of materials and the upper limits of table 1. the regulations define a "hardship or infeasibility exemption" which may be invoked if a) there is a lack of commercially available material necessary to comply with the regulations; b) the cost of complying with the regulations is disproportionate to the overall cost of the project; or c) compliance with the regulations would impair the integrity of a recognized historic building. since passage of the marin county ordinance, there has been considerable interest from other county and city governments in considering similar regulations, particularly in the state of california. 4.4. the buy clean california act the california state legislature passed the "buy clean california act" in 2017 [10]. this legislation requires that, for state-funded projects, information about the global warming potential of the materials used in the project must be reported. the act also mandates that embodied carbon of construction materials must be considered as one of the decision factors when selecting materials, contractors and materials suppliers. 607 andrew warren taylor acta polytechnica ctu proceedings minimum specified concrete strength f ′ca maximum ordinary portland cement contentb maximum embodied carbon, from epd [mpa] [kg/m3] ! kg (co2e) /m3 " ≤ 17.2 215 260 20.7 243 289 27.6 271 313 34.4 298 338 41.3 315 356 48.2 352 394 > 48.2 390 433 ≤ 20.7 light weight 304 578 27.6 light weight 339 626 34.4 light weight 373 675 a for concrete strength between the stated values, use linear interpolation to determine cement and/or embodied carbon limits. b portland cement of any type meeting the requirements of astm c150 [8]. table 1. marin county cement and embodied carbon limits. certain construction materials are addressed by the act, but at this time cement and concrete are not among them. steel reinforcing bars are currently covered by the act, and it is reasonable to expect that cement and concrete will eventually be included. any construction material that is listed under the act is assigned an upper limit on global warming potential; subsequently, when that material is purchased by the state a version of that material must be supplied that meets the global warming potential limit. because the state of california has a large economic and political influence in the western united states, it is likely that other western states will consider adopting similar "buy clean" rules for state-funded construction projects. 4.5. new york state legislation on september 12, 2019, a proposal was introduced to the new york state legislature that would require the use of concrete with low embodied carbon on new york state projects (i.e. all state funded projects that incorporate concrete as a construction material). this legislation would also promote the use of concrete that implements co2 capture technologies. in addition, the proposal introduced a tax credit system that would promote the preparation of epds by allowing concrete suppliers to deduct the cost of preparing epds from their tax bill. although this proposed legislation was not adopted, a revised version will likely be re-introduced to the new york state legislature in the near future. 5. discussion given the recent surge of interest in the global warming potential of construction materials in general, and cement and concrete in particular, it is incumbent upon codes and standards writing organizations to develop frameworks for quantifying and reporting the global warming potential of construction projects. because each construction material is unique in terms of its manufacture, transport, and implementation, it naturally falls upon the professional and industrial organizations related to each material to develop such guidelines. complementary to the work of professional and industrial organizations to understand the global warming potential of specific construction materials, there is the work of local, state, and national government organizations to develop guidelines and regulations that will lead to reductions in greenhouse gas emissions. these government agencies must balance the technical, economic, and political factors that influence the creation of environmental policies. it is the job of structural and materials engineers to ensure that government agencies have the technical data and analytical tools necessary to make informed decisions about the environmental impacts of construction materials. this is why the american concrete institute has embarked on a multi-faceted program to study, quantify, and develop engineering solutions that will reduce the environmental impacts of cement and concrete. section 3.4 above describes aci’s current activities in this area. 6. conclusions this paper provides an overview of current activities in the united states related to building codes and regulations governing the environmental impacts of cement and concrete. this is a societal issue that is evolving rapidly, not only in the united states, but worldwide. some of the pending regulations described in this paper may soon be enacted into law, and other new initiatives will soon be introduced in local, state, and national jurisdictions. in an effort to support this trend towards improved control of greenhouse gas emissions, various groups in the united states are developing metrics for quantifying 608 vol. 33/2022 sustainable concrete codes in the usa the global warming potential of construction materials. this presents an opportunity for technical organizations in the united states, such as the american concrete institute, to provide leadership in matters related to quantifying and reporting the global warming potential of cement and concrete. acknowledgements the author wishes to thank the american concrete institute (aci) for travel funding provided through the aci ambassador program. references [1] so 14025:2006. environmental labels and declarations type iii environmental declarations principles and procedures, geneva: iso, 2006. [2] n. post, j. gonchar. embodied carbon tool premieres engineering news record, new york: bnp media ii, .p 8-9, 2019. [3] architecture 2030, 2020. http://architecture2030.org. [4] aci committee 130. 130r-19: report on the role of materials in sustainable concrete construction, farmington hills: american concrete institute, 2019. https://cecollection2.files.wordpress.com/2020/ 05/130r-19-report-on-the-role-of-materials-i n-sustainable-concrete-construction.pdf. [5] aci committee itg-10. report on alternative cements, farmington hills: american concrete institute, 2018. https://www.concrete.org/portals /0/files/pdf/previews/itg-10.1r-18_preview.pdf. [6] aci committee itg-10. practitioner’s guide for alternative cements, farmington hills: american concrete institute, 2018. https://www.concrete.org/portals/0/files/pdf/pr eviews/itg-10r-18_preview.pdf. [7] state of washington. house bill 2412: an act relating to creating the buy clean washington act; and adding a new chapter to title 39 rcw, olympia: washington state house of representatives 65th legislature, 2018. [8] american society for testing and materials international 2019 astm c150/c150m 19a: standard specification for portland cement, west conshohocken: astm international, 2019. [9] marin county california. section ii: marin county code chapter 19.07 added to marin county code title 19, marin: marin county board of supervisors, p. 1-6, 2019. [10] state of california. public contract code division 2, part 1, chapter 3, article 5: buy clean california act, sacramento: california state legislature, 2017. 609 http://architecture2030.org https://cecollection2.files.wordpress.com/2020/05/130r-19-report-on-the-role-of-materials-in-sustainable-concrete-construction.pdf https://www.concrete.org/portals/0/files/pdf/previews/itg-10.1r-18_preview.pdf https://www.concrete.org/portals/0/files/pdf/previews/itg-10r-18_preview.pdf doi:10.14311/app.2021.30.0030 acta polytechnica ctu proceedings 30:30–35, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app mechanical and chemical activation of blast furnace slag and its influence on the mechanical properties of the resulting cement paste jan horych∗, pavel tesárek, zdeněk prošek czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jan.horych@fsv.cvut.cz abstract. the usage of waste materials is a very important global topic. the large amount of waste everywhere in the world needs to be processed or disposed. landfilling is not an option anymore, because of european legislation and restrictions. a lot of studies are trying to develop new options or possibilities of using waste materials. this research is trying to find a way to process blast furnace slag. a high-speed mill was used for the mechanical activation. chemical activation was used as the next step of activation. there are many materials that could be used, but in this study we used slaked lime and water-glass. slaked lime had a positive effect on mechanical properties. samples had higher compressive strength but the effect was limited only for 5 wt. %. another used material was water-glass, but in this case, there was a significant negative effect. compressive strength and flexural strength were significantly reduced. keywords: activation, blast furnace slag, cement replacement, slaked lime, water-glass. 1. introduction the objective of this study is to find a way to replace cement from concrete with useless waste materials. according to the trend of recycling we tried to find an option to use old and highly carbonated blast furnace slag as a partial substitute for portland cement. concrete is a trendy material thanks to its variability and reliability. cement is the main and the most important ingredient of concrete, but it is also the most expensive ingredient from natural resources. the production itself is highly economically and ecologically demanding. an issue with cement is not a new topic. for example, bellmann f. and stark j. [1] focused on chemical activation of blast furnace slag by adding calcium hydroxide and soluble calcium salts such as calcium chloride, calcium bromide, calcium nitrate, calcium formate and calcium acetate. early compressive strength was measured on mortar samples. compressive strength was increased by adding calcium hydroxide, calcium carbonate and calcium acetate from 3 mpa to 25 mpa after 7 days. after the 28 days period the final compressive strength was increased from 36 mpa to 53 mpa. another research from yum w. et al. [2] investigated the influence of calcium formate on the strength and microstructure of the ground granulated blast furnace slag activated with calcium oxide. the next study shows the effect of different slag characteristics on compressive strength in slag system with calcium oxide activation written by jeong y. et al. in 2016 [3]. four kinds of ground granulated blast furnace slags from different sources were collected to investigate changes in final mechanical properties. each of the slags developed significantly different compressive strength regardless of seemingly similar characteristics of the slags. the values of the final compressive strengths were spread from 25 mpa up to 52 mpa in 28 days. the reaction products were different as well. the goal of the study by wang y. et al. [4] was to determine the effect of wet grinding on the mechanical properties of blast furnace slag and slag-cement mixtures. testing was focused on different grinding times from 10 min to 50 min in 10 min intervals. the performance of well ground samples of slag indicated the effectiveness of wet grinding. the 50 min grinding could decrease the average particle size to 2.95 µm. other effects were high reactivity, improvement of mechanical properties and reduction of porosity. research of escalante-garcía j. et al. [5] observed a mechanically activated granulated blast furnace slag. samples were composed in a range of 50 to 95 wt. % to replace clinker in portland slag cement. structural properties were determined using x-ray diffraction, electron microscopy, thermogravimetry and differential thermal analysis. up to the 70 wt. % replacement was found an improvement of compressive strength in the first and 28th day. microstructural changes of the slag caused by mechanical activation provided enhanced reactivity and the improvement of cement strength. the paper from sobolev k. [6] examined the effect of grinding on high-performance cement containing granulated blast furnace slag. samples were produced in high volumes of blast furnace slag. testing results of mortars based on modified cement improved compressive strength up to 91.7 mpa. final mechanical properties were increased by 62 % over the reference sample. 30 http://dx.doi.org/10.14311/app.2021.30.0030 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 mechanical and chemical activation of blast furnace slag 0 10 20 30 40 50 60 70 80 90 100 0.25 1 4 16 64 256 cu m ul at iv e am ou nt [% ] particle size [µm] slag o not ground slag a cem i 42.5 r figure 1. particle size distribution of used materials. mixtures cement[wt. %] slag [wt. %] type of slag slaked lime [wt. %] w/b [-] bulk density [kg/m3] c100 100.0 0.40 2020 ± 10 o30/a0 70.0 30.0 o 0.40 1978 ± 29 o30/av5 67.5 27.5 o 5 0.40 1958 ± 17 o30/av10 65.0 25.0 o 10 0.40 1940 ± 9 a30/a0 70.0 30.0 a 0.40 1992 ± 28 a30/av5 67.5 27.5 a 5 0.40 1998 ± 22 a30/av10 65.0 25.0 a 10 0.40 1942 ± 25 table 1. composition of individual samples with slaked lime. 2. materials and samples during the research, portland cement was still a basic component. specifically, it was portland cement cem i 42.5 r from radotín. another component was blast furnace slag. the source of slag was a foundry industry complex in kladno. it was a waste from slag aggregate production by recycling. used slag was approximately 100 years old air-cooled and carbonated blast furnace slag stored on a heap. at first, waste materials were not immediately ready for another use. it needed to be activated. as mentioned, mechanical activation was used on the base of previous study of prošek z. [7]. collaboration with lavaris ltd. was used again. the main goal was to find a material useable as a chemical activator of waste materials. dimensions of all produced samples were 40 × 40 × 160 mm. they were stored in water for 28 days. during this time, we tested every sample with non-destructive methods to determine dynamic young’s modulus of elasticity and dynamic shear modulus. at the end of this period, samples were destroyed with a destructive bending test and compression test to confirm final mechanical properties. slaked lime was used for the first experiment. in this test we also used two samples of slag with the different grinding capacity to find an ideal option and investigate the synergy of mechanical and chemical activation. figure 1 shows the difference between individual samples of the slags. the slag with no grinding was called slag o. as we can see, slag a was ground and the particle size distribution changed significantly and is similar to the curve of the reference cement cem i 42.5 r. samples were composed of changing shares of portland cement 65 to 100 wt. %, 25 to 30 wt. % of blast furnace slag and 0 to 10 wt. % of slaked lime (table 1). water binder ratio (w/b) was always set on 0.4. in the second phase water-glass was used based on sodium and potassium. the composition of samples was similar to the previous tests. there was 65 to 100 wt. % of portland cement, 25 to 30 wt. % of blast furnace slag and 0 to 10 wt. % of the activator (table 2). based on the results of the first phase, testing samples were produced using only slag a. 3. measurement methods several methods were used through testing to investigate mechanical properties. first of them was a nondestructive resonance method. in period of 28 days dynamic young’s modulus of elasticity and dynamic shear modulus were determined every week. a set of instruments and devices was used to measure natural frequency of all samples in longitudinal, transverse 31 j. horych, p. tesárek, z. prošek acta polytechnica ctu proceedings 0 5 10 15 20 25 30 0 7 14 21 28 y ou ng 's d yn am ic m od ul us [g p a] time [days] c100 a 30/a0 a 30/av5 a 30/av10 o 30/a0 o 30/av5 o 30/av10 figure 2. development of dynamic young’s modulus of the samples with slaked lime. mixtures cement[wt. %] slag [wt. %] sodium silicate [wt. %] potassium silicate [wt. %] bulk density [kg/m3] c100 100.0 2020 ± 10 a30/a0 70.0 30.0 1969 ± 36 a30/an5 67.5 27.5 5 1963 ± 27 a30/an10 65.0 25.0 10 1899 ± 10 a30/ak5 67.5 27.5 5 1945 ± 8 a30/ak10 65.0 25.0 10 1899 ± 13 table 2. composition of individual samples with water-glass. and torsional oscillation. the supports of samples were soft and elastic pads and had to be at node lines of basic natural mode shapes. impact hammer type 8206 was used to excite vibration of the samples. acceleration transducer type brüel & kjær type 4519-003 measured the response of the samples. both of devices are connected to the measurement station 3560-b-x12. it’s purpose is to record the excitation and response signals. the station transforms received signals to the frequency domain using the fast fourier transform. computer with specific software called pulse labshop then creates a frequency response function. the dynamic modulus of elasticity and the dynamic shear modulus were calculated according to astm 1876-01 and čsn 73 1372 [8, 9]. the measurement of the dynamic modulus of elasticity and the dynamic shear modulus was taken from the work of j. topič [10]. the flexural strength was measured at the end of the 28 days period by using three-point bending test on samples measuring 40 × 40 × 160 mm. samples were broken approximately in half (40 × 40 × ∼80 mm). the loading through the testing was displacement controlled at a constant rate of 1 mm/min. the distance between the supports was 100 mm. then these halves were used to determine compressive strength by uniaxial compression test. uniaxial compression testing was displacement controlled at a rate of 5 mm/min [11, 12]. 4. results and discussion as mentioned through testing, the development of young’s dynamic modulus and dynamic shear modulus was measured for all samples. in the first phase of the experiment, slaked lime was used to activate blast furnace slag. figure 2 shows the development of dynamic young’s modulus. first we can see a difference approximately 4 gpa between a reference sample with 100 wt. % of portland cement and the rest of the samples. it is caused by the replacement of portland cement with blast furnace slag. another slight difference, we can see between samples with a lower amount of slaked lime. samples a 30/av0, a 30/av5 and o 30/av0 had a little bit higher values of dynamic young’s modulus and dynamic shear modulus (figure 3). the highest progress of both modules was registered in first 7 days. later testing from day 7 to day 21 showed lower improvement. in the last week of testing, values were almost the same. results of non-destructive testing suggest there was an influence of mechanical activation. slag a had a slightly higher both dynamic young’s modulus and dynamic shear modulus. also, the effect of slaked lime as an activator was limited. only 5 wt. % of 32 vol. 30/2021 mechanical and chemical activation of blast furnace slag 0 2 4 6 8 10 12 14 0 7 14 21 28 d yn am ic s he ar m od ul us [g p a] time [days] c100 a 30/a0 a 30/av5 a 30/av10 o 30/a0 o 30/av5 o 30/av10 figure 3. development of dynamic shear modulus of the samples with slaked lime. figure 4. flexural strength of the samples with slaked lime. figure 5. compressive strength of the samples with slaked lime. slaked lime had a positive effect on final properties. at the end of the first phase, all samples were tested with destructive methods. the final flexural strength is shown in figure 4. we can see a significant positive effect of blast furnace slag on flexural strength. in some cases, values were almost tripled. in general, slag has a much slower reaction and produces less heat. it needs time to react appropriately. for these reasons, the concrete based on slag is used for a long time. the influence of slaked lime was registered on the samples a 30/av5 and o 30/av5. lower values of flexural strength suggest a more complex chemical reaction. figure 5 presents compressive strength and results overall confirmed previous conclusions. mechanical activation had a positive effect on the final properties. limited value of 5 wt. % of slaked lime supported a chemical reaction and helped to increase compressive strength. slag a was more finally ground and grains were more exposed. with slaked lime, we achieved the value 56 mpa. the positive effect of slaked lime is perceptible. in the second phase, water-glass based on sodium and potassium was used to investigate its influence on final mechanical properties. similar to the previous tests through 28 days period, dynamic young‘s modulus and dynamic shear modulus were measured every week. as we can see in figure 6 and figure 7, the final values are more spread. unfortunately, we can see a big difference between each samples. the reference sample had significantly higher values. the difference was approximately 5 gpa. the second highest value of dynamic young’s modulus and dynamic shear modulus had a sample a 30/a0 without any activator. by adding a water-glass as an activator, the final values were even more reduced. with more activator, the value drop was more obvious. the lowest values had samples with the highest amount of waterglass, regardless of sodium or potassium. the highest progress was also measured in the first 7 days. later development was almost completely stopped and dynamic shear modulus was even decreased in the last week. water-glass seemed unusable and negative effect was significant. the destructive tests confirmed the results of the resonance tests. flexural strength of the samples 33 j. horych, p. tesárek, z. prošek acta polytechnica ctu proceedings 0 5 10 15 20 25 30 0 7 14 21 28 y ou ng 's d yn am ic m od ul us [g p a] time [days] c100 a30/a0 a30/an5 a30/an10 a30/ak5 a30/ak10 figure 6. development of dynamic young’s modulus of the samples with water-glass. 0 2 4 6 8 10 12 14 0 7 14 21 28 d yn am ic s he ar m od ul us [g p a] time [days] c100 a30/a0 a30/an5 a30/an10 a30/ak5 a30/ak10 figure 7. development of dynamic shear modulus of the samples with water-glass. figure 8. flexural strength of the samples with water-glass. with water-glass was reduced. with a higher amount of water-glass, the flexural strength drop was even more significant. figure 8 shows the final results and values of flexural strength. the sample a 30/ak10 with 10 wt. % of water-glass based on potassium had figure 9. compressive strength of the samples with water-glass. almost the same flexural strength as the reference sample c 100 with 100 wt. % of portland cement. the negative effect of water glass almost removed the influence of the blast furnace slag. 34 vol. 30/2021 mechanical and chemical activation of blast furnace slag compressive strength shown on figure 9 proved water-glass does not work as an activator. the samples with water-glass had significantly reduced compressive strength. it seemed water-glass partially prevents the chemical reaction. the cement and blast furnace slag grains could not create a proper connection. the sample a 30/ak10 had the compressive strength reduced almost on half compared to the sample a 30/a0 and less than third compared to the sample c 100 only with portland cement. unfortunately, the result of the whole second phase was a complete failure. water-glass can not be used as the activator. 5. conclusions this study investigated the influence of mechanical and chemical activation of the blast furnace slag to support a possibility of cement replacement up to the level of 35 wt. %. another goal was to find an ideal chemical activator and determine synergy between mechanical and chemical activation. based on the results, it can be concluded that: • blast furnace slag had a positive effect on flexural strength. • highly carbonated blast furnace slag can be used as a binder to replace portland cement. • mechanical and chemical activation created a synergy and supported the hydration process. • slaked lime worked as an activator but at a limited amount of 5 wt. %. • water-glass did not work as an activator and even significantly reduced the influence of blast furnace slag regardless of sodium or potassium. acknowledgements this paper was financially supported by czech technical university in prague under no. sgs project sgs19/148/ohk1/3t/11. the authors also thank lavaris ltd. for the supplied materials. references [1] f. bellmann, j. stark. activation of blast furnace slag by a new method. cement and concrete research 39(8):644–650, 2009. doi:10.1016/j.cemconres.2009.05.012. [2] w. s. yum, j. i. suh, d. jeon, j. e. oh. strength enhancement of cao-activated slag system through addition of calcium formate as a new auxiliary activator. cement and concrete composites 109:103572, 2020. doi:10.1016/j.cemconcomp.2020.103572. [3] y. jeong, h. park, y. jun, et al. influence of slag characteristics on strength development and reaction products in a cao-activated slag system. cement and concrete composites 72:155–167, 2016. doi:10.1016/j.cemconcomp.2016.06.005. [4] y. wang, x. he, y. su, et al. efficiency of wet-grinding on the mechano-chemical activation of granulated blast furnace slag (gbfs). construction and building materials 199:185–193, 2019. doi:10.1016/j.conbuildmat.2018.11.245. [5] j. i. escalante-garcía, a. f. fuentes, a. gorokhovsky, et al. hydration products and reactivity of blast-furnace slag activated by various alkalis. journal of the american ceramic society 86(12):2148–2153, 2003. doi:10.1111/j.1151-2916.2003.tb03623.x. [6] k. sobolev. mechano-chemical modification of cement with high volumes of blast furnace slag. cement and concrete composites 27(7-8):848–853, 2005. doi:10.1016/j.cemconcomp.2005.03.010. [7] j. trejbal, z. prošek, p. tesárek, j. valentin. activation of blast furnace slag and its usage in practise. týden výzkumu a inovací pro praxi 2019 2019. [8] astm 1876-01. standard test method for dynamic young’s modulus, shear modulus, and poisson’s ratio by impulse excitation of vibration, annual book of astm standards, american society for testing and materials 2006. [9] čsn 73 1372. in non-destructive testing of concrete – testing of concrete by resonance method, czech office for standarts, metrology and testing. 2012. [10] j. topič, z. prošek. properties and microstructure of cement paste including recycled concrete powder. acta polytechnica 57(1):49–57, 2017. [11] čsn en 12390 (731302). in testing hardened concrete, czech office for standarts, metrology and testing. 2013. [12] j. kot’átková, j. zatloukal, p. reiterman, k. kolář. concrete and cement composites used for radioactive waste deposition. journal of environmental radioactivity 178:147–155, 2017. 35 https://doi.org/10.1016/j.cemconres.2009.05.012 https://doi.org/10.1016/j.cemconcomp.2020.103572 https://doi.org/10.1016/j.cemconcomp.2016.06.005 https://doi.org/10.1016/j.conbuildmat.2018.11.245 https://doi.org/10.1111/j.1151-2916.2003.tb03623.x https://doi.org/10.1016/j.cemconcomp.2005.03.010 doi:10.14311/app.2021.30.0001 acta polytechnica ctu proceedings 30:1–6, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app size effect on the ultimate drying shrinkage of cement mortar: 1-year experiment and numerical modeling lenka dohnalová∗, petr havlásek, vít šmilauer, pavel reiterman, vendula davidová czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: lenka.dohnalova@fsv.cvut.cz abstract. the magnitude and time evolution of shrinkage are influenced by numerous factors which are implemented in the design codes often in a different way. the time-dependent behavior of concrete in structures sensitive to creep and shrinkage should be verified by means of short-term laboratory measurements. extrapolation of drying shrinkage from short-term measurements is an illposed problem. the process is extremely slow but can be accelerated by reducing the specimen size. the knowledge of the size-effect on drying shrinkage is a necessity to establish the transition from the laboratory to the structural size. in the literature, the experimental data on such size-effect are insufficient. for this reason a new experiment was developed to study this phenomenon on small-scale specimens made of cement mortar and the results from the first year are summarized in this paper. the measured data are validated by coupled fem hygro-mechanical simulations. keywords: concrete, drying, modeling, shrinkage, size effect. 1. introduction creep and shrinkage of concrete are complex timedependent processes whose physical origin remains still fully unresolved. the importance of both phenomena and the need for their correct prediction grew with the boom of pre-stressed concrete technology which allowed to design slender structures with larger spans and which allowed to utilize concrete potential economically. at the same time, such challenging structures are generally more sensitive to the timedependent deformations in concrete. the magnitude and time evolution of shrinkage is influenced by numerous factors (ranging from concrete composition to ambient conditions) which are not approached consistently among current prediction models and design codes for concrete. to name a few, the subtle interplay between the development of drying shrinkage and drying creep (investigated e.g. in [1]) can determine whether the durability of concrete structure will be jeopardized by tensile cracking or the shrinkage-induced tensile stresses are relieved sufficiently fast by increased compliance. a disagreement among the different models prevails also in the question of the size-effect on the ultimate value of drying shrinkage [2] which is addressed in this paper. the resulting discrepancies between the design codes give rise to many uncertainties which can drastically affect the requirements on the design of shrinkageand creep-sensitive structures. the prediction of the behavior of such structures can be improved by laboratory measurements on specimens prepared from the same concrete composition. this approach can be successfully utilized to specify the estimated concrete creep based on short-time measurements. however, the direct applicability to drying shrinkage is rather limited because the extrapolation presents an ill-posed problem. the characteristic time and the ultimate shrinkage cannot be determined uniquely until the rate plotted in log-scale significantly decreases [3]. in the case of standard laboratory specimens 70 × 70 × 285 mm or 100 × 100 × 500 mm, a subtle decrease in shrinkage rate (in log scale) occurs after approx. 100 days while the almost ultimate shrinkage strain is reached after approx. 1000 days of drying [4]. the size-effect on drying shrinkage is in a certain way introduced in majority of design codes or models for time-dependent behavior of concrete. the origin of the size-effect can be sought in nonuniform distribution of relative humidity over the cross-section which gives rise to non-uniform shrinkage strains. the shrinkage strains are internally restrained which produces highly nonlinear distribution of self-equilibrated stresses. owing to the presence of these stresses in the early stages of drying, concrete is extremely prone to surface cracking. the generated stresses become redistributed as the concrete dries and their magnitude is gradually relieved by concrete creep. this complicated interplay causes that the value of ultimate drying shrinkage decreases with increasing size of the concrete member. if the size-effect on the ultimate shrinkage is known, the time-demanding experiments can be significantly accelerated when the dimensions of a spec1 http://dx.doi.org/10.14311/app.2021.30.0001 http://ojs.cvut.cz/ojs/index.php/app l. dohnalová, p. havlásek, v. šmilauer et al. acta polytechnica ctu proceedings figure 1. images of the experimental setup: frames for automated (left) and manual (right) readings. imen become reduced. the rate of drying and thus (drying) shrinkage scales with the square of the effective thickness d defined in b3 model [5] as 2v /s where v is the volume of the specimen and s is the surface exposed to drying. for example in the case of an infinite prism with square cross-section, d coincides with half of its thickness. in this paper, parameter d is used to express the specimen size. the comprehensive creep and shrinkage database assembled at the northwestern university [6] comprises a remarkable number of experimental results gathered from the entire world. however, the promising content of this database becomes rather limited when only a specific phenomenon is to be identified. only a small number of experimental studies examined drying shrinkage of specimens with different sizes prepared from the same concrete mixture and subject to identical conditions. this selection needs to be reduced even further because in many cases the experiment was terminated prematurely before the ultimate shrinkage was reached. a recent paper by the present authors [2] attempted to evaluate the aforementioned size-effect based on the data available in the literature and the aforementioned database. the experiments unanimously indicate that the ultimate drying shrinkage decreases with specimen size, but this relationship cannot be evaluated more accurately owing to insufficient relevant experimental data. the lacking data stimulated the present authors to design and execute a new narrow-oriented experiment focused directly on the size-effect on drying shrinkage. this paper presents and summarizes the measured data collected within the first year of the experiment and compares the strain development in time with the results obtained from coupled hygromechanical finite element simulations. 2. materials and methods 2.1. experimental setup the proposed setup for the shrinkage and moisture loss measurement comprised a large variety of prismatic specimens with uniform length 400 mm but with cross-section ranging from 20 × 20 mm (d = 10 mm) to 100 × 100 mm (d = 50 mm). in order to guarantee material homogeneity and to obtain representative data even on the smallest specimens, cement mortar with 2 mm maximum aggregate size was preferred at the expense of concrete. the composition of the mortar mix had standard sand-to-cement ratio 3:1 by mass. standard cen siliceous sand and the blended binder cem ii 32.5 b-s with slag content 29% were used. the typical water-to-cement ratio was reduced from w/c = 0.5 to 0.45 to match the behavior of mortar in a typical structural concrete. compressive strength and young’s modulus measured at the age of 33 days on standard cylinders 100 × 200 mm were 33.9 mpa and 40.9 gpa, respectively. the slender prisms were cast into a custom-built steel formwork with plastic bottom. the goal was to resemble curing in sealed conditions. for this reason the top was covered with 1 mm thick pe sheet followed by a damp cloth. the following day the specimens were transported to a sealed container with henv = 96%. the specimens were demolded at the age of 31 days when the drying began. the ends of the prisms were equipped by metal sheets with a central dent to provide suitable surface for the measurement as well as moisture sealing. the reference measurement was done once the glue had hardened. altogether 11 mortar prisms with one companion steel specimen for temperature compensation were placed in a special measuring frame, see fig. 1 left. each specimen rested on a bolt with round head. axial shortening was measured at the top using linear transducer with internal spring return mmr1012 (effective stroke 12.7 mm, total resistance 10 kω). data logging occurred automatically, the setup did not allow to measure the moisture loss due to drying. for this reason, a second group consisting of 15 specimens was measured in a more conventional approach using stiff measuring frame equipped with digital indicator sylvac with 1 µm precision, see fig. 1 right. this setup allowed to determine also the evaporated water which was recorded using precision balance scales; kern 572-39 (range 4.2 kg, readout 0.01 g) and t-scale (range 15 kg, readout 0.1 g) for sample with cross-section 100 × 100 mm. minimum difference between the coefficient of thermal expansion of steel and mortar allowed relative reading, circumventing temperature compensation. the experiment was executed in a laboratory without humidity and temperature control. in order to take the inevitable fluctuations into consideration, both temperature and relative humidity were recorded. the frequency corresponded to the automatic and manual measurements. 2.2. numerical modeling the consistency of the measured data was verified using one-way coupled hygro-mechanical simulations 2 vol. 30/2021 size effect on the drying shrinkage 0 200 400 600 800 1000 1200 1400 400 1 10 100 0.3 0.4 0.5 0.6 0.7 0.8 0.9 humidity sh ri nk ag e, ε sh [ × 1 06 ] a m bi en t h um id ity , h en v drying duration, t t0 [day] d = 10 mm d = 13 mm* d = 15 mm d = 20 mm* d = 20 mm d = 25 mm d = 30 mm d = 50 mm 0 200 400 600 800 1000 1200 400 1 10 100 0.3 0.4 0.5 0.6 0.7 0.8 0.9 humidity sh ri nk ag e, ε sh [ × 1 06 ] a m bi en t h um id ity , h en v drying duration, t t0 [day] d = 10 mm d = 13 mm* d = 15 mm d = 20 mm* d = 20 mm d = 25 mm d = 30 mm d = 50 mm figure 2. evolution of shrinkage measured automatically (left) and manually (right). the shaded area corresponds to the experimental range in the case more identical specimens. asterisk denotes specimens with rectangular crosssection (square cross-section otherwise). 0 200 400 600 800 1000 1200 1400 400 1 10 100 0.3 0.4 0.5 0.6 0.7 0.8 0.9 humidity sh ri nk ag e, ε sh [ × 1 06 ] a m bi en t h um id ity , h en v drying duration, t t0 [day] d = 10 mm d = 15 mm d = 20 mm d = 25 mm d = 30 mm d = 50 mm figure 3. comparison of the average shrinkage strain measured automatically (solid lines) and manually (symbols). the two techniques gave good agreement during the first 6 months but afterwards started to diverge which was attributed to the deformation of the plywood back of the shrinkage frame for automated measurements. in open-source finite element package oofem [7]. concrete drying is, in the present study, described by a widely recognized model proposed by bažant and najjar [8]. under certain simplified assumptions, the governing equation for the diffusion of water vapor reads ∂h ∂t = ∇ · (c(h)∇h) (1) where ∇h is the gradient of relative humidity and c(h) is the humidity-dependent diffusivity. this dependence is for the cementitious materials highly nonlinear and can be approximated as c(h) = c1 ! "α0 + 1 − α0 1 + # 1−h 1−hc $n % & (2) where c1 is the maximum diffusivity at h = 1, α0 determines the ratio between minimum diffusivity at h = 0 and c1, and parameters hc and n describe the relative humidity threshold and the rate of the transition. the ambient relative humidity is prescribed using mixed boundary condition which relates the humidity flux jh with the humidity difference at the boundary via surface factor f . in each time step, the moisture transport is followed by the structural sub-problem which utilizes the computed field of relative humidity. the mechanical problem uses constitutive model based on the microprestress-solidification (mps) theory [9]. under constant temperature and sealed conditions, the mps model reduces to the b3 model [5]. however, changes in relative humidity give rise not only to volume changes (shrinkage or swelling), but also to additional creep (pickett effect). the model has been modified [10] to minimize the size-effect on drying creep which is controlled by parameter k3. shrinkage strain and relative humidity are linearly linked via their rates, ε̇sh = kshḣ (3) where ksh is a material parameter usually treated as a humidityand age-independent constant. the mps model is extended with the cohesive crack model and fixed orientation of cracks to consider material softening if the tensile strength is reached. however, the formed cracks do not influence the humidity transport, allowing one-way coupling scheme. the computational model was calibrated in several subsequent steps. the basic creep compliance, which is in the case of the b3 model described by 4 parameters q1 − q4, was determined from empirical prediction formulae using compressive strength and concrete composition as input parameters. afterwards, parameters controlling short-term creep q1 and q2 were adjusted to match the results of the 1-hour creep experiment. the value of the drying creep parameter, k3 = 26 was identified in the previous analysis of bryant and vadhanavikkit experiment 3 l. dohnalová, p. havlásek, v. šmilauer et al. acta polytechnica ctu proceedings b3 model mps model moisture diffusivity + mixed b.c. q1 q2 q3 q4 k3 ft gf ksh c1 α0 hc n f [×10−6/mpa] [ ] [mpa] [n/m] [ ] [mm2/day] [ ] [ ] [ ] [mm/day] 15.5 150 2.08 9.41 26 3.0 100 0.0016 30 0.15 0.7 12 0.3 table 1. summary of material parameters [11]. the response is only very slightly sensitive to the value of tensile strength ft and fracture energy gf , which were set to typical values for concrete (see tab. 1 which summarizes the values of all basic parameters used in simulations). the remaining parameter of mps model ksh, and the parameters of the transport model were calibrated to achieve reasonable agreement with the shrinkage results of the performed experiment. the aim was to achieve the best possible match for all prism sizes. the main emphasis was placed on the initial development of shrinkage (approx. first 6 months of drying) when both groups of the specimens show similar development of shrinkage. later, when there is no longer a match between the specimen groups with a different measurement method, the aim was to achieve the best possible match with the results of the manually measured prisms. for the purpose of identification of material parameters, the evolution of ambient relative humidity was defined by a piecewise-linear function according to the measured values. in addition, the analysis was repeated for a partially adjusted development of ambient humidity (beginning according to the experiment, after 91 days of drying assumed humidity fixed at henv = 60%) achieve final values of shrinkage and to demonstrate more clearly the size-effect on drying shrinkage predicted by the model. 3. results and discussion the development of shrinkage in time is for both automated and manual measurement techniques shown in fig. 2. in the case of more identical specimens, the mean response is shown in solid line while the shaded area corresponds to the measured experimental range. the obvious fluctuations in shrinkage strain are caused by oscillating value of ambient relative humidity which is shown in dashed black line. the average responses obtained with the two techniques are compared in fig. 3. except of the largest specimen, the measurements demonstrate very good agreement during the first 6 months of drying. sudden decrease in relative humidity from 60% to 40%, which occurred afterwards, caused rapid acceleration of drying accompanied by increase in shrinkage rate. this decrease also led to separation of the experimental trends determined by the two methods; in all cases the shrinkage strain measured automatically exceeded the manual readings. the reason for this was not clarified even by thorough analysis of the measured data. still, the most probable cause seems to 0 200 400 600 800 1000 400 1 10 100 sh ri nk ag e, ε sh [ × 1 06 ] drying duration, t t0 [day] d = 10 mm d = 15 mm d = 20 mm d = 25 mm d = 30 mm d = 50 mm figure 4. comparison of the experimental measurements (symbols) with the results obtained from fem simulations (solid lines). be connected to the frame for automated measurement. the measuring frame was attempted to be sufficiently stiff for the current measurement. since the frame serves only as a support for the specimens and is loaded merely by their self weight, the selected material was 26 mm thick plywood for concrete molds. however, this material turned out to be insufficiently stable under variable relative humidity of the ambient environment since non-negligible volume changes were observed. these changes could not be completely eliminated by incorporating the measurement of the steel standard primarily intended for temperature compensation. the difference between the shrinkage of geometrically identical specimens determined by the automated and manual measurement technique was not uniform and was found to be, in a certain way, influenced by their position, namely by the distance from the largest specimen placed at one end (see fig. 1, the weight of the prism labeled “1” reached almost 10 kg). the joint effect of these two factors seems to be the most probable explanation. therefore, the data from the manual measurements are preferred in the later stage of the experiment (solid steel structure of the measuring frame insensitive to level of relative humidity) despite the intervals between the individual measurements were gradually increasing. as was stated above, the manual measurements allowed to determine not only the axial shrinkage but also the evolution of weight loss due to drying. to the best knowledge of the authors, such data are not available in this extent in any of the preceding experiments related to the size effect on drying shrinkage. 4 vol. 30/2021 size effect on the drying shrinkage 0 10 20 30 40 50 60 70 80 400 1 10 100 0.3 0.4 0.5 0.6 0.7 0.8 0.9 humidity w ei gh t l os s [k g/ m 3 ] a m bi en t h um id ity , h en v duration of drying, t t0 [day] d = 10 mm d = 15 mm d = 20 mm d = 25 mm d = 30 mm d = 50 mm 0 10 20 30 40 50 60 70 80 400 1 10 100 w ei gh t l os s [k g/ m 3 ] drying duration, t t0 [day] d = 10 mm d = 15 mm d = 20 mm d = 25 mm d = 30 mm d = 50 mm figure 5. experimentally measured moisture loss (left) and its comparison to fem analysis (right). after 100 days of drying the model exhibits excessive sensitivity to changes in henv. 0.4 0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 0 50 100 150 200 250 300 350 400 reference size, d = 10 mm si ze e ff ec t drying duration t t0 [day] d = 13 mm* d = 15 mm d = 20 mm* d = 20 mm d = 25 mm d = 30 mm 0.4 0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 0 50 100 150 200 250 300 350 400 0.3 0.4 0.5 0.6 0.7 0.8 reference size, d = 10 mm si ze e ff ec t a m bi en t h um id ity , h en v drying duration t t0 [day] d = 15 mm d = 20 mm d = 25 mm d = 30 mm humidity figure 6. time development of the experimentally measured (left) and computed (right) size-effect on drying shrinkage. the data points correspond to the average values obtained from manual measurement. the ambient relative humidity recorded in the experiment resulted into the fem response shown in solid lines while the modified history of humidity (henv = 60% since 91 days of drying) led to the results shown in dashed lines. yet such data provide better understanding of the time development of the drying process. the measured evolution of moisture loss normalized per volume of the specimen is shown in the left part of fig. 5. the sensitivity to changes in ambient relative humidity which occurred in the later stage of drying is in the case of moisture loss not so pronounced as in the case of drying shrinkage. in the finite element simulations the driving force for the moisture transport is the relative humidity, therefore the moisture loss needs to be backcalculated. the value of moisture capacity (slope of the desorption isotherm, for simplicity assumed as linear) which gave the best agreement with the experimental data as shown in right part of fig. 5 was k = 120 kg/m3. in the simulations, the development of ambient relative humidity corresponded with the experimental measurements. the agreement with the experimental data is satisfactory only in the first 100 days of drying; afterwards, the model overestimates the rate of the weight loss or gain caused by changes in henv. the discrepancy in the initial value of weight loss stems from the linear desorption isotherm which is not realistic in the case of high relative humidity. the ongoing changes in henv did not allow the specimens to reach their ultimate value of drying shrinkage. for this reason the size effect on drying shrinkage is here evaluated within the entire duration of the experiment and is related to the thinnest specimen with cross-section 20 × 20 mm (d = 10 mm). naturally, the strongest size-effect can be observed shortly after the onset of drying when the difference between the moisture loss of specimens with different size is most remarkable. on the other hand with the increasing duration the size-effect is slowly diminishing until the equilibrium with the ambient conditions is reached. the expected evolution of the size-effect on drying shrinkage is confirmed by the experimental measure5 l. dohnalová, p. havlásek, v. šmilauer et al. acta polytechnica ctu proceedings ments as shown in the left part of fig. 6. even though the ultimate values of drying shrinkage have not been reach so far, the size-effect has become almost stable and there is no indication that is should vanish completely. the processed data are in correct order, the size-effect increases with size. with respect to the reference specimen (d = 10 mm) the size effect is within 20% of that value. as was shown in fig. 4, the experimentally measured evolution of shrinkage was captured by the numerical model very realistically up to ≈ 180 days of drying. however, the mismatch beyond this time causes that the numerically computed size-effect significantly deviates from the experiment as demonstrated by the smooth solid lines in the right part of fig. 6. in order to arrive at equilibrated values of drying shrinkage, the simulation was repeated with a modified history of henv, which was fixed at 60% after 91 days of drying. under such conditions the predicted size-effect becomes almost negligible within one year of drying. 4. conclusions this paper analyzed and presented the results from the first year of a pilot experiment developed to investigate the size-effect on drying shrinkage. the primary motivation of this research was to verify whether the shrinkage measurements could be accelerated using smaller specimens than is the current practice. the findings can be summarized as follows. • the difficulties encountered in the first run of the experiment provided the research team with a valuable feedback for its future improvement. the experimental technique itself is feasible, cheap and capable to deliver good-quality data. a stiff steel frame, automatic monitoring of ambient conditions are a necessity, while the currently used monitoring of moisture loss is sufficient. • the measured data were in expected order and exhibited very little scatter among specimens with the same dimensions. this indicates good homogeneity of the material in the specimens despite their small dimensions. moreover, the consistency of the data was verified by a computational model and the agreement was promising. • the size-effect on drying shrinkage cannot be properly evaluated unless the experiment is conducted at constant relative humidity, which was not the present case. despite this deficiency, in the currently examined size range a non-negligible sizeeffect on drying shrinkage was registered. • the constitutive model used in structural analysis needs to be improved as it strongly underestimates the size-effect on drying shrinkage. acknowledgements the authors gratefully acknowledge financial support from the czech science foundation (ga čr), project number 19-20666s. references [1] l. dohnalová, p. havlásek. comparison of drying shrinkage and drying creep kinetics in concrete. acta polytechnica ctu proceedings 15:12–19, 2018. [2] l. dohnalová, p. havlásek. size effect on the ultimate drying shrinkage of concrete experimental evidence and engineering practice. acta polytechnica ctu proceedings 26:13–18, 2020. [3] z. bažant, m. jirásek. creep and hygrothermal effects in concrete structures. springer netherlands, 2018. doi:10.1007/978-94-024-1138-6. [4] c. aguilar. estudio del comportamiento y desarrollo de una metodologia de prediccion de la retraccion hidraulica de hormigones, phd. thesis (in spanish). pontificia universidad catolica de chile. escuela de ingenieria, 2005. [5] z. bažant, s. baweja. creep and shrinkage prediction model for analysis and design of concrete structures: model b3. adam neville symposium: creep and shrinkage structural design effects 2000. [6] m. hubler, r. wendner, z. bažant. comprehensive database for concrete creep and shrinkage: analysis and recommendations for testing and recording. aci 112:547–558, 2015. [7] b. patzák. oofem home page, 2000. http://www.oofem.org. [8] z. p. bažant, l. j. najjar. nonlinear water diffusion in nonsaturated concrete. materials and structures 5:3–20, 1972. doi:10.1007/bf02479073. [9] z. p. bažant, a. p. hauggaard, s. baweja, f. j. ulm. microprestress solidification theory for concrete creep. i: aging and drying effects. journal of engineering mechanics 123:1188–1194, 1997. [10] z. bažant, p. havlásek, m. jirásek. microprestress-solidification theory: modeling of size effect on drying creep. in n. bicanic, h. mang, g. meschke, r. de borst (eds.), computational modelling of concrete structures, pp. 749–758. crc press/balkema, eh leiden, the netherlands, 2014. [11] a. h. bryant, c. vadhanavikkit. creep, shrinkage-size, and age at loading effects. aci materials journal 84:117–123, 1987. 6 https://doi.org/10.1007/978-94-024-1138-6 https://doi.org/10.1007/bf02479073 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2021.31.0058 acta polytechnica ctu proceedings 31:58–60, 2021 © 2021 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague maintaining the competitiveness of the czech railway industry jan vogl∗, mária jánešová czech technical university in prague, faculty of transportation sciences, konviktská 20, 110 00 prague 1, czech republic ∗ corresponding author: vogljan@fd.cvut.cz abstract. the development of foreign trade plays an important role in the czech economy. the important foreign trade commodity item is machinery and transport equipment (sitc 7). railway vehicles is a part of this category, other means of transport (sitc 79). in order to maintain competitiveness and economic growth, it is necessary for the railway industry to intensify the innovation and the introducing of new technologies into production. one of the available options is a wider implementation of life cycle cost analysis, which enables to activate innovative efforts, as well as to improve technical and performance data. keywords: export, railway vehicles, lcc analysis, innovation. 1. introduction the czech republic is a traditional export industrial country. therefore, the development of foreign trade plays an important role in our economy. last year, exports from the czech republic grew by 3.6% to czk 4 563.3 billion. the largest volume of czech exports has been realized in the long term within the european union countries, such as germany, france, great britain and poland. in 2019, the share of exports to eu countries reached 83.6%. in terms of commodity structure (sitc), machinery and transport equipment (sitc 7) are traditionally dominant. their share in total exports in 2019 was about 59%. compared to 2018, it grew by 5.1% [1]. figure 1. cross-border movements of goods by commodities, year 2019, source: czso [1]. 2. description of product group railway vehicles is part of lager category, other transport equipment (sitc 79), which consists of: • 791 railway vehicles (including hovertrains) and associated equipment • 792 aircraft and associated equipment • 793 ships, boats and floating structures if we focus on the export development in the commodity item railway vehicles and related equipment (railway vehicles (including hovertrains) and associated equipment) we find out that in 2019 this item reached the value of czk 20.5 billion. the largest volume of czech exports of the mentioned commodity goes to europe (approx. 88%). these are, for example, deliveries of locomotives, metro trains, trams and components [1]. the published analysis of the association of the czech railway industry (acri) regarding the development of the czech railway industry in 2012-2019 shows that 2015 was a turning point for the companies, which are the members of this association. in 2015, they have reached the highest turnover of 91 billion czk. for example the total turnover in 2019 is about 16.5% lower in comparison with 2015, as well as the total number of employees is lower by 6.5%. due to the global competition in foreign markets, czech railway industry is not able to reach the successful year of 2015 up to now [2]. 3. innovation and lcc analysis in order to remain competitive on foreign markets, we need to increase the innovative potential of our products. the czech republic will not keep its significant competitive advantages of cheap labour for long time. according to a study by the european commission entitled "country-specific recommendations 2020 research and innovation analysis", the czech republic lags behind in the field of research and innovation within the eu. productivity growth is mostly driven by large foreign companies, while domestic companies are lagging behind. the same applies to the growth of expenditure on research and development [3]. 58 https://doi.org/10.14311/app.2021.31.0058 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 31/2021 maintaining the competitiveness of the czech railway industry figure 2. exports of railway vehicles (including hovertrains) and associated equipment, source: acri [2]. the annual analysis of the european commission "european innovation scoreboard 2020" showed that the czech republic was, according to the results of 27 indicators included in the moderate innovators [4]. figure 3. performance of eu member states’ innovation systems, source: eu [4]. innovation of the offered products is a key task for maintaining the competitiveness of manufacturing companies in the market. this can be achieved through innovation and close cooperation between companies and research institutions in research and development. the introduction of new technologies and innovations is important for the further successful economic growth of czech industry. efficient use of life cycle cost (lcc) analysis contributes to this growth. the aim of lcc analysis is to determine the total cost of a product throughout its life cycle [5]. the results of my research among leading manufacturer of railway vehicles have shown that lcc analysis is one of those methods that has the ability to help increase innovation effects, the quality of the offered product and reduces overall costs. furthermore, lcc analysis plays a significant contribution in tenders to select the railway vehicles. 4. conclusion the increasing use of the lcc analysis should be an important incentive for manufacturers to improve the reliability and quality of their products. the implementation of the lcc analysis will lead to innovative efforts and thus increase the competitiveness of the company. in the rail sector, where railway vehicles has been in operation for more than 30 years [6], operators can, together with manufacturers, optimize the total cost of ownership, which accounts for the largest share of the total lcc cost by railway vehicles [5]. based on university research conducted in 2019-2020 with leading manufacturer of railway vehicles, it turned out that the use of lcc analysis not only reduces overall costs, but also activates innovation. as the czech industry is dependent on the export of its products to foreign markets, it is necessary not only in the railway industry, which has a two-hundred-year tradition, to capture new trends in the autonomous control, alternative drives or intelligent transport systems [7]. this requires intensive and consistent cooperation with the government, research institutions, high-tech companies, but also other sectors such as energy and information and communication technologies. one of the approved government strategic framework plans is the innovation strategy of the czech republic 2019-2030, which aims to help move the czech republic among the most innovative countries in europe within a few years. if we manage to capture new world trends in this sector, the czech republic will continue the economical grow and the improving of the quality of offered products and services to our citizens. references [1] český statistický úřad [čsu]. pohyb zboží přes hranice. in: český statistický úřad. databáze zahraničního obchodu v přeshraničním pojetí [online]. ©2020. available in: https: //apl.czso.cz/pll/stazo/stazo.stazo?jazyk=en. [2] analýza asociace podniků českého železničního průmyslu. in: acri [online]. praha: asociace podniků českého železničního průmyslu, 13. 5. 2020. available in: https://acri.cz/2020/05/24/firmy-zeleznicnihoprumyslu-jsou-stabilnimi-tahouny-ceskeekonomiky-dokazuji-to-jejich-hospodarskevysledky-za-rok-2019/. [3] country-specific recommendations 2020 – research and innovation analysis. [online]. european commission. brussel, 2020. available in: https://rio.jrc.ec.europa.eu/library/2020country-specific-recommendations-research-andinnovation-analysis. [4] european innovation scoreboard 2020. [online]. european commission. brussel, 2020. available in: https://ec.europa.eu/commission/presscorner/ detail/en/qanda_20_115. [5] čsn en 60300-3-3. management spolehlivosti – část 3-3: pokyny použití – analýza nákladů životního cyklu. praha: český normalizační institut, 2004. 60 s. [6] kontrolní závěr z kontrolní akce 17/35: pořízení a obnova železničních kolejových vozidel. in: nejvyšší kontrolní úřad [online]. praha. available in: https://www.nku.cz/assets/kon-zavery/k17035.pdf. 59 https://apl.czso.cz/pll/stazo/stazo.stazo?jazyk=en https://apl.czso.cz/pll/stazo/stazo.stazo?jazyk=en https://acri.cz/2020/05/24/firmy-zeleznicniho-prumyslu-jsou-stabilnimi-tahouny-ceske-ekonomiky-dokazuji-to-jejich-hospodarske-vysledky-za-rok-2019/ https://acri.cz/2020/05/24/firmy-zeleznicniho-prumyslu-jsou-stabilnimi-tahouny-ceske-ekonomiky-dokazuji-to-jejich-hospodarske-vysledky-za-rok-2019/ https://acri.cz/2020/05/24/firmy-zeleznicniho-prumyslu-jsou-stabilnimi-tahouny-ceske-ekonomiky-dokazuji-to-jejich-hospodarske-vysledky-za-rok-2019/ https://acri.cz/2020/05/24/firmy-zeleznicniho-prumyslu-jsou-stabilnimi-tahouny-ceske-ekonomiky-dokazuji-to-jejich-hospodarske-vysledky-za-rok-2019/ https://rio.jrc.ec.europa.eu/library/2020-country-specific-recommendations-research-and-innovation-analysis https://rio.jrc.ec.europa.eu/library/2020-country-specific-recommendations-research-and-innovation-analysis https://rio.jrc.ec.europa.eu/library/2020-country-specific-recommendations-research-and-innovation-analysis https://ec.europa.eu/commission/presscorner/detail/en/qanda_20_115 https://ec.europa.eu/commission/presscorner/detail/en/qanda_20_115 https://www.nku.cz/assets/kon-zavery/k17035.pdf jan vogl, mária jánešová acta polytechnica ctu proceedings [7] memorandum o budoucnosti českého železničního průmyslu. in: acri [online]. praha, 2019. available in: https://acri.cz/2019/04/12/memorandum-obudoucnosti-ceskeho-zeleznicniho-prumyslu/. 60 https://acri.cz/2019/04/12/memorandum-o-budoucnosti-ceskeho-zeleznicniho-prumyslu/ https://acri.cz/2019/04/12/memorandum-o-budoucnosti-ceskeho-zeleznicniho-prumyslu/ acta polytechnica ctu proceedings 31:58–60, 2021 1 introduction 2 description of product group 3 innovation and lcc analysis 4 conclusion references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0144 acta polytechnica ctu proceedings 26:144–148, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app pre-processing of additive manufacturing input files for numerical simulation jan voříšek∗, bořek patzák czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jan.vorisek@fsv.cvut.cz abstract. in this contribution, we present the concept of a 3d printer software emulator facilitating the creation of a spatial finite element mesh suitable for the printing process simulation. the concept is based on gradual processing of a native 3d printer input file (in a g-code format). this file contains a complete description of manufacturing process consisting of series of individual commands interpreted by a printer. the effect of each command needs to be precisely evaluated to obtain the position of the printer head and the volume of the deposited material in any given time. the calculation is performed in the same way as in the marlin printer firmware using the trapezoidal motion curves and a command buffer. to represent a computational model, a discrete voxel model with variable edge length and time discretization is used. the volume of deposited material is calculated for each voxel as a function of time. the resulting model is suitable for numerical analysis of the printing process. keywords: additive manufacturing, 3d printing, 3d printer emulator. 1. introduction additive manufacturing emerged as an innovative manufacturing process in the field of rapid prototyping. it is based on incremental deposition or sintering of the material in thin layers. wide range of materials, including plastic, metal or even concrete, can be used to produce spatial objects of complex geometries and shapes (see figure 1). the objects can be designed using currently available cad software. printing a prototype is often more affordable than making one by hand, and certainly less costly than contracting a manufacturer to do it by conventional manufacturing, especially in the fields where low quantities of parts with a high degree of customization and complexity play the critical role. an extensive overview of the additive manufacturing methods is available in [1]. additive manufacturing consists of several independent sub-processes. everything begins with a 3d model which, in order to be printed, must be processed using specialised software called slicer. the slicing process produces a sequence of commands to control the movement and operation of the machine stored in a g-code file line by line. this file describes the complete processing (it is directly interpreted by the printer) and is used as an input for the presented pre-processing tool. visualization of the movement commands can be seen in figure 2 and figure 4. in recent years, there has been an increasing interest in the simulation of the additive manufacturing process. parts produced primarily with extrusion of molten thermoplastics often do not reach designed dimensions due to the cracking and warping induced by large temperature gradients [2, 3]. such issues are also figure 1. visualization of the geometry of a printed object in cad. figure 2. visualization of the movement commands contained in the input file (single layer). moves with material extrusion are drawn with blue color, moves without extrusion with black color. 144 https://doi.org/10.14311/app.2020.26.0144 https://ojs.cvut.cz/ojs/index.php/app vol. 26/2020 pre-processing of additive manufacturing input files present in conventional manufacturing processes like injection moulding where accurate numerical simulation allows to mitigate such issues by optimising the object geometry or the production process itself [4]. numerical simulation of the manufacturing process is indispensable to produce tailored, defect-free and optimised products. due to the complexity of the process, the coupled transient thermo-mechanical analysis is needed for reliable predictions. some research of the additive manufacturing has been already conducted using the finite difference method [3]. in this contribution we assume that the finite element method (fem) will be used to solve the problem. our goal is to develop a software tool able to automatically prepare inputs for the fem model (domain discretization, boundary conditions, etc.) straight from the additive manufacturing input files. not only should our tool generate a suitable 3d mesh, but it should also provide the history of the infill percentage of the single mesh elements. by knowing the infill percentage, we will be able to consider the two-phase characteristics of the voxels (thermoplastics and air). the process is, in a sense, an inverse process of slicing as it produces the simplified (voxelized) version of the original object. several g-code processing utilities are available online [5, 6]. these tools provide print time estimation and filament consumption. however, they only consider the printer moves as vectors, and thus constructing the voxel model with the material extrusion history is not possible. our solution extends the idea of the movement commands into the 3d space considering each of them as a block of material with a finite volume. 2. the model finite element model can generally consist of elements of various shapes. we have chosen a discrete voxel model with variable edge length along each axis due to its simplicity. assuming all the edge lengths equal leaves us with a simple cube. this simplification will allow the precomputation of shape function values and their derivatives as they are same for all the elements. on top of that, such a hexahedral element is commonly used in the fem software packages. 3. g-code pre-processing the first step into the simulation of additive manufacturing is to pre-process the raw input files. these files contain a line-separated set of instructions for the 3d printer to follow. a complete overview of all the available g-codes is available online [7]. only a small subset of the commands, which controls the deposition of the material, is considered. processing of the printer input file begins with determining the duration of individual commands and particularly of printer head moves. this step is performed only once and results can be reused to generate simulation inputs with different resolution. 3.1. duration of moves it is essential to capture the real printing speed in the sequence of point-to-point moves. the target printer speed is defined in the g-code file for each move, yet the actual motion profile is to be decided by the individual printer. constant acceleration is the most used planning strategy resulting in a trapezoidal velocity profile due to its mathematical simplicity and limited computing power of the 3d printers microprocessors. the printer stores a limited number of moves, usually 16, in a buffer and recalculates the acceleration profile allowing the printer to be able to stop after the last move, if needed. the typical printing speed of one print layer can be observed in figure 3. maximal accelerations and velocities must be respected in the direction of each axis. these limits can be taken from the 3d printer specification or the source code of the printer controller firmware (marlin, reprap). duration of the movements can be determined using the calculated velocity and known distance. by gradually summing up the duration of individual commands, we can obtain the current time for any given print head position. this information is essential for numerical simulation as we can reverse this procedure to calculate the position of the print head for any given time (time step). figure 3. visualization of the printer head speed during processing of a single layer assuming the trapezoidal velocity profile. 3.2. reconstructing the extrusion width extrusion width is essential for calculating the intersection of the extrusion volume and the individual voxels to determine the amount of material deposition inside each voxel. it is difficult to define the exact shape of the extrusion cross-section. therefore slicers use various simplified shapes for their internal calculations, most notably a rectangle with curved sides (fig. 6) or a simple rectangle [8]. 145 jan voříšek, bořek patzák acta polytechnica ctu proceedings figure 4. g-code visualization of the first layer of a 20 × 20 × 20 mm cube. figure 5. activated voxel elements of a first layer of a 20×20×20 mm cube including volume assignment. extrusion volume shape is automatically set in the slicing software based on the nozzle diameter or manually input by the user. however, the information is not present in the exported g-code file, and so it must be recalculated using known variables. to make the calculation slicer-independent, we assume a rectangular cross-section of the deposition. the extruded volume can be calculated as vin = wihili (1) where wi is extrusion width, hi is extrusion volume height and li is extrusion volume length (see figure 6). it must be equal to the volume of material coming into the extruder given as vin = πd2 4 ∆e (2) where d is nozzle diameter and ∆e is length of extruded material. the unknown extrusion width is then expressed as figure 6. extrusion width geometry used to produce g-code in slic3r software [8] wi = πd2 4hili ∆e (3) the simplified calculation produces a slightly narrower extrusion width than the original slic3r value assuming oval-shaped sides of the cross-section. the difference is negligible and will be not be taken into account in any further calculations. for slicers assuming rectangular extrusion cross-sections, the calculated width will match the original value. extrusion width calculation can be omitted for additive manufacturing processes based on the laser beam sintering of material such as stereolitography. in such case the width is defined by the beam diameter. 4. voxel activation and filling the voxel-activation and the extruded volume assignment is performed for each combination of time-step length and voxel-size. see figure 5 for an example of a single processed print layer. 4.1. the principle single depositions, represented by block geometry, are to be decomposed (projected) into the individual voxel contributions. calculating the intersection of two generally oriented blocks would be very computationheavy. fortunately, the deposition of the material occurs in horizontal planes, and so the intersection can be decomposed into two much simpler processes. the first process is the so called horizontal decomposition, where extrusion volume (represented as arbitrarily oriented rectangle is decomposed to individual voxel column contributions (rasterization). in the second step (vertical decomposition), the extruded volume is distributed to individual voxels in a column according to the height and position of extruded material volume. these two subprocesses together provide history of extruded material volume in each voxel. 146 vol. 26/2020 pre-processing of additive manufacturing input files figure 7. activation sequence of the individual voxel elements. the colour of each voxel corresponds to its infill percentage (vof). the centre of the print consists mostly of filled elements shown in red colour while blue elements near the surface of the printed part are almost empty. 4.2. optimization the bounding box of the depositions is calculated to save computer resources, so only the voxels contained within the bounding box are considered for the intersection calculation. furthermore, it is possible to use bresenham’s line algorithm to only mark the affected voxels by doing only integer math. 4.3. issues some issues emerged after implementing the abovedescribed process. slicers usually produce overlapping extrusions (see figure 8) to improve the bonding of the parallel extrusions. the resulting infill history contained over-filled voxels (volume over 100%). the issue is also caused by considering the depositions as non-continuous (see figure 9). the second issue can be eliminated by considering the intersection of successive depositions. the overlapping volume is rotated by 180 degrees and moved into the affected voxels as can be seen in figure 9. figure 8. overlap of the two concurrent depositions (overlapping volume marked with red). figure 9. overlap of the two non-straight depositions (overlapping volume marked with red). the red volume must be redistributed into the green area. the voxel volume is checked after each volume difference assignment, and the surplus volume (over 100%; if applicable) is redistributed into the eight horizontal neighbouring elements to keep the maximum volume of each voxel at 100%. 5. conclusions prototype g-code pre-processor and voxel-model generator were implemented in javascript based on the previously formulated methodology. several realworld files were processed using the tool. example of complex geometry can be seen in figure 7. so far, it appears that it is possible to create voxelbased geometry. the developed tool enables to observe the time sequence of element activation immediately after the input file analysis. for example the model in the figure 7 was processed on a generic consumer laptop in one thread using a 1 × 1 × 1 m voxel grid in 2.6 seconds. it is desirable to perform 3d numerical analysis of the additive manufacturing process using the knowledge of element activation sequence and the vof distribution over time. 6. future work our main goal is to create a robust, user-friendly tool for the numerical analysis of additive manufacturing and its optimization. the tool should be able to take the g-code input file, process it to obtain mesh with element infill history and finally, run fem analysis and post-process the results. this paper solves the preprocessing part as the first step in an overall complex process. for proper numerical simulation, it is desired to identify the thermal, mechanical, as well as processrelated properties of the materials involved. we foresee the formulation of suitable constitutive material models and development of suitable adaptive fe solver. furthermore, we also expect to conduct lab-scale measurements to determine constitutive parameters and develop and calibrate suitable material models. the idea of gradual voxel activation could be extended from to the activation of generally shaped elements of a finite element mesh. the elements generally do not need to form any regular grid. therefore current advanced mesh generators could be used to discretize printed object. such elements can be gradually activated and filled in a same principle as the voxels. however, we expect the intersection of the extrusion block and a general element to be very computationally expensive. 147 jan voříšek, bořek patzák acta polytechnica ctu proceedings acknowledgements the financial support of this research by the grant agency of the czech technical university in prague (sgs project no. sgs19/032/ohk1/1t/11) is gratefully acknowledged. references [1] h. bikas, p. stavropoulos, g. chryssolouris. additive manufacturing methods and modelling approaches: a critical review. the international journal of advanced manufacturing technology 83(1-4):389–405, 2015. doi:10.1007/s00170-015-7576-2. [2] y. zhang, k. chou. a parametric study of part distortions in fused deposition modelling using three-dimensional finite element analysis. proceedings of the institution of mechanical engineers, part b: journal of engineering manufacture 222(8):959–968, 2008. doi:10.1243/09544054jem990. [3] t. stockman, j. a. schneider, b. walker, j. s. carpenter. a 3d finite difference thermal model tailored for additive manufacturing. jom 71(3):1117–1126, 2019. doi:10.1007/s11837-019-03338-6. [4] r. johansson, d. konijnendijk. injection moulding simulation — a finite element approach to analyse thethermodynamics in an im tool. master’s thesis, queensland university of technology, 2007. [5] a. ustyantsev. gcodevisualizer a web-based visual gcode viewer and analyzer. https://github.com/hudbrog/gcodeviewer. accessed: 2019-07-11. [6] gcodeanalyser.com. g-code analyser. http://www.gcodeanalyser.com/. accessed: 2019-07-11. [7] reprap. g-code — reprap. https://reprap.org/wiki/g-code. accessed: 2019-07-10. [8] g. hodgson, a. ranellucci, j. moe. slic3r manual — flow math. https://manual.slic3r.org/advanced/flow-math. accessed: 2019-07-15. 148 https://doi.org/10.1007/s00170-015-7576-2 https://doi.org/10.1243/09544054jem990 https://doi.org/10.1007/s11837-019-03338-6 https://github.com/hudbrog/gcodeviewer http://www.gcodeanalyser.com/ https://reprap.org/wiki/g-code https://manual.slic3r.org/advanced/flow-math acta polytechnica ctu proceedings 26:145–149, 2020 1 introduction 2 the model 3 g-code pre-processing 3.1 duration of moves 3.2 reconstructing the extrusion width 4 voxel activation and filling 4.1 the principle 4.2 optimization 4.3 issues 5 conclusions 6 future work acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0107 acta polytechnica ctu proceedings 26:107–111, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app mupif workflow editor and automatic code generator stanislav šulc∗, vít šmilauer, bořek patzák czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: stanislav.sulc@fsv.cvut.cz abstract. integrating applications or codes into mupif api model enables easy integration of such apis into any workflow representing complex multiphysical simulation. this concept of mupif also enables automatic code generation of the computational code for given workflow structure. this article describes a ’workflow generator’ tool for the code generation together with ’workflow editor’ graphical interface for interactive definition of the workflow structure and the inner data dependencies. the usage is explained on a thermo-mechanical simulation. keywords: mupif, workflow editor, code generator. 1. introduction complex physical tasks often need integration of very different models or solvers. since there exist models solving the particular tasks, it is a very effective approach to connect the existing software tools together to solve the complex tasks. a mupif platform [1] is a python [2] tool for such interconnection of several models together to compute coupled or linked multiphysical simulation. note that in our terminology a model can be some software, application or code. each model is covered in a python api derived from mupif class model, which has the following functions: __init__() # creates an instance of the class model. initialize() # initializes the model instance for a specific usecase. solvestep() # solves a computational step with defined length. terminate() # cleans all resources of the model instance. set() # sets inputs of the model, e.g. mupif field or property. get() # returns outputs of the model, e.g. mupif field or property. getcriticaltimestep() # returns maximum time step length these seven functions suffice for full control of a basic computation. the communication between models is realized with exchanging instances of mupif classes field or property. these instances can be obtained from a model by calling its get() method with appropriate parameters, and can be sent to a model by calling its set() method. both field and property instances keep their values, but they also have attributes describing their type, units and some additional info for unique identification. then, a property can be sent to a model and the model decides what to do with it. this concept creates environment for easy interconnection of even completely different codes, which can also be based on different programming languages. the value of a property is a tuple of any value types. a field has a tuple of nodal or cell values for each entity, according to the field type, while each value can be a tuple. a field is always linked with a mupif mesh, which gives the field space definition, interpolation and portability to tasks with different discretization. several models together create a workflow, which can be executed or added into another workflow, as it can behave like a model from outer scope, see fig. 1. the mupif class workflow is derived from class model and is extended with the following function: solve() # calls solvestep() until the target time is reached. which solves the whole simulation. a simplified python code of a simulation with two models can look like this: tm = thermal_model() mm = mechanical_model() tm.initialize(’input_t.in’, ...) mm.initialize(’input_m.in’, ...) while time<targettime: tm.solvestep(...) tf = tm.get(...) mm.set(tf) mm.solvestep(...) ... tm.terminate() mm.terminate() 2. mupif workflow generator the concept of mupif, especially the unified structure of the apis allows us to automate generation of the computational steering code when the simulation 107 https://doi.org/10.14311/app.2020.26.0107 https://ojs.cvut.cz/ojs/index.php/app s. šulc, v. šmilauer, b. patzák acta polytechnica ctu proceedings figure 1. example of a workflow structure. figure 2. workflow diagram of a thermo-mechanical simulation. workflow structure is defined. each api contains list of inputs and outputs which can be interconnected according to the data flow in the chain of models. this list also contains detailed description of each input or output. the mupif workflow generator project [3] provides python code environment to define the hierarchical structure of the simulation. it contains classes of blocks, where each block represents an item in the workflow diagram. there are blocks for following entities: mupif model, mupif property, mupif physical quantity, time loop and the whole workflow. there are also some other blocks and also a possibility of implementing a new block by user. workflow and time loop blocks can contain a sequence of another blocks. all blocks can have input and output data slots. in case of model block they represent the model inputs and outputs. in case of a property or field the block has one output slot representing its value. two data slots (output and input) can be connected, which represents the data exchange between models or setting of the input parameters. when all the blocks and the data dependencies are defined, the code can be generated. the mupif workflow code can be generated in two ways. the first option is that the workflow structure defines function solvestep(). in this case we call it a class workflow. the second option is that the workflow has a timeloop in itself and the structure defines the function solve(). in this case we call it an execution workflow. in case of a class workflow there is a possibility of setting some inputs or outputs, just like a model instance has some. it is realized using external data slots of the workflow. these slots are connected to some input or output slots inside the workflow, which transfers data from or to the outer scope, see this logic in fig. 1. this is further explained in section 4. the execution workflow cannot have such slots because the generated code behaves like a script to be run, not like a module to set some parameters and call its specific functions as in case of class workflow. 3. mupif workflow editor mupif workflow editor [4] is a graphical user interface of the workflow generator for easy definition of the simulation structure according to its moda [5] representation. each item in the workflow is represented as a block with list of input slots on the left and list of output slots on the right. these slots can be interactively linked to define the dependencies among 108 vol. 26/2020 mupif workflow editor and automatic code generator particular models, see fig. 4. the input slots can also be linked to some constant input data represented with blocks of mupif constant property, field or physical quantity. there is also a block to define a time-loop, a block for including custom code or ifelse block. when all the necessary dependencies are defined, the python code of the simulation can be generated and saved or executed instantly. the workflow editor provides saving of the workflow state into a json file to save the work and continue later. at the present time, it takes a part in the eu horizon 2020 project and thus it contains a list of all available involved models inside, but any model based on mupif model or workflow can be loaded into this tool from a python module. it automatically creates the structure and visual representation with the available data slots according to the inputs and outputs specification. 4. example of use the usage of the workflow editor and generator is described on a simple nonstationary thermo-mechanical simulation. this task is one of the mupif examples and can also be found in the workfloweditor github repository [4]. the geometry and boundary conditions are depicted in fig. 3. figure 3. schema of a thermo-mechanical task. the material parameters are: conductivity 1.0 w.m−1.k−1, capacity 1.0 j.kg−1.k−1, density 1.0 kg.m−2, young’s modulus e=30×109 pa, ν=0.25 and coefficient of thermal expansion 1.2×10−6 k−1. the mechanical response is caused only by the thermal expansion, there is no force load. the moda workflow representation of the task is depicted in fig. 2. 4.1. execution workflow we begin with the option of defining structure of the whole simulation in one execution workflow. the data flow from fig. 1 is realized in the workflow editor as shown in fig. 4. the thermal solver solves a nonstationary 2d thermal task on a rectangular domain. the boundary conditions on all edges can be set to the solver, see the thermal nonstat model inputs in fig. 4. the left and bottom edge temperature inputs are linked to the output of a block representing mupif constant temperature property of zero value. the top edge temperature is linked analogically to a block representing constant temperature of 10 °c. figure 4. execution workflow of the thermomechanical simulation. the output temperature field slot of the thermal task is linked to the input temperature field slot of the mechanical task, which yields in the following code: mm.set(tm.get(mupif.fid_temperature, ts.gettime())) 109 s. šulc, v. šmilauer, b. patzák acta polytechnica ctu proceedings where tm is the thermal model instance, mm is the mechanical model instance and ts is the computational time step. the geometry of both thermal and mechanical task is defined via input file for each task. the block blocktimeloop represents the simulation time loop and contains four model blocks. the last two blocks represent model named field export to vtk, which is a simple tool for saving given mupif field into a file. it must have been integrated into the mupif class model to enable such usage in the workflow editor. the mupif temperature and displacement fields are exported into files within each computational timestep. the first three blocks of the workflow represent mupif time physical quantity and define the run of the simulation time loop. 4.2. class + execution workflow we continue with the second option, which begins with definition of a class workflow with the thermal and mechanical models, see fig. 5. figure 5. class workflow of the thermo-mechanical simulation. the data slots are connected analogically, but the top edge temperature slot is linked to the external input data slot and thus this data requirement will be satisfied from outer scope. note that this input of the workflow is called ’external input data slot’ although inside the workflow its value is represented with an output data slot to be linked with another block’s input data slot. in case of external outputs it works analogically, see the two external output data slots, creating the outputs of this workflow. now we generate the class code and save it into a file. it contains a definition of a class derived from class workflow. then we create another workflow and insert a blocktimeloop into it, see fig. 6. figure 6. execution workflow of the thermomechanical simulation with thermo-mechanical workflow used as a model. this block contains the thermo-mechanical workflow as a model and illustrates the encapsulation of a complex simulation into a workflow which provides easier usage from outer scope. 4.3. task solution and results both approaches from section 4.1 and section 4.2 yield into an execution workflow for which we can generate the python code. it can be saved into a file or executed instantly from the workflow editor. the execution will produce the vtk [6] files, which can be easily visualized, see figs. 7 and 8. 110 vol. 26/2020 mupif workflow editor and automatic code generator figure 7. resulting temperature at time=10 s. figure 8. resulting displacement at time=10 s. 5. conclusions this article describes a graphical software tool for user-friendly definition of a multiphysical simulation structure and the following generating of a python code of the simulation. the usage of the workflow editor tool was explained on a thermo-mechanical task in two modes the first one is the whole task defined in one execution workflow. the second one is defining a class workflow to be imported into an execution workflow as a model. the example is available in the workflow editor github repository. in the future, a web-based version of the workflow editor will be created to provide easier access to this tool without any installation. acknowledgements we gratefully acknowledge financial support from eu horizon 2020 project, contract number: 721105. we also gratefully acknowledge the financial support from czech technical university in prague, grant number sgs19/032/ohk1/1t/11. references [1] mupif https://github.com/mupif/mupif, 2019. [2] python https://www.python.org/. [3] mupif workflow generator project repository. https://github.com/mupif/workflowgenerator, 2019. [4] mupif workflow editor project repository. https://github.com/mupif/workfloweditor, 2019. [5] the european materials modelling council moda, homepage. https://emmc.info/moda/. [6] vtk https://vtk.org/. 111 acta polytechnica ctu proceedings 26:107–111, 2020 1 introduction 2 mupif workflow generator 3 mupif workflow editor 4 example of use 4.1 execution workflow 4.2 class + execution workflow 4.3 task solution and results 5 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2021.31.0027 acta polytechnica ctu proceedings 31:27–29, 2021 © 2021 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague high-speed rail – highway 2.0? pavel krupík czech technical university in prague, faculty of civil engineering, department of construction management and economics, thákurova 7/2077, 166 29 prague 6 dejvice, czech republic correspondence: pavel.krupik@fsv.cvut.cz abstract. high-speed rail lines should make it possible to fully replace air transport over the medium distances of europe (about 500 km). in the czech republic, they are beginning to design. estimates suggest that in five years, the first sections could begin to physically originate and within ten years, they will begin to ride on them. is it possible that the scenario and the continuous extension of deadlines accompanying the completion of the backbone network will be repeated? this article discusses the options from the project management perspective and the idea of construction 4.0 to prevent this. keywords: high-speed rail, project management, predictive project life cycle, stakeholders, construction 4.0. 1. introduction the supreme audit office made two basic conclusions after taking an inventory of the preparation and construction of motorways between 2013 and 2017. the better news is: motorway prices have fallen by 55 percent on average compared to the years 2008 to 2012. the bad one: the preparation of constructions is getting longer and the basic motorway network will not be completed even in 2050 [1]. in 2018, 833 kilometers of the planned 2,073 kilometers were to be built to complete the backbone network of motorways. to achieve this, the road and motorway directorate would have to open approximately 25 kilometers of motorway annually. but the preparation is not accelerating, on the contrary. compared to the previous inspection in 2012, the phase from obtaining a favorable opinion of the eia to the issuance of a building permit extended by four years to an average of 13 years [1]. is the state failing? is it possible that the situation will repeat itself in the preparation and construction of high-speed lines? the initial situation is as follows. regulation (eu) no 1315/2013 of the european parliament and of the council identifies the main principles for the development of road, rail, inland waterway, maritime and multimodal infrastructure. rail transport is further divided into two parts of ten-t [2]: • the comprehensive network aims to connect all european nuts ii regions to the ten-t network, with an indicative completion date of 2050. network continuity has become a basic requirement. great emphasis has also been placed on cross-border sections. for the territory of the czech republic, the ten-t regulation identifies the main directions of high-speed rail development for further examination. • the core network is a subset of the global network and has been established by the european commission on the basis of a uniform european methodology. according to this methodology, the main network includes all the links of the neighboring so-called primary nodes (simply the capitals of the states and agglomerations over 1 million inhabitants). the second condition is a guarantee for the completion of the project by the end of 2030, provided that its economic efficiency is demonstrated. ministry of transport of the czech republic released program for the development of fast railway connections in the czech republic, signed by resolution of the government of the czech republic no. 389 of 22 may 2017 [3, 4]. by 2030, three sections of high-speed lines are expected to be realized within the ten-t main network, namely the new prague – lovosice / litomerice construction including the exit from the prague junction, the major reconstruction of the brno prerov line and modernization and capacity increase of the brno – breclav line. the new construction of the prague – brno high-speed line in its entire length was not enforced into the ten-t main network, as it is not feasible until the end of 2030, given the current process delays in the construction preparation period. the ministry of transport reached this conclusion on the basis of an assessment of experience from abroad, where implementation of similarly large projects from the beginning of preparation to completion usually takes 15 – 20 years. however, the potential of this connection within the czech republic is considerable and this section should be significantly preferred in the preparation, see below. however, the construction time limit is neither the design preparation, the actual construction work, nor the provision of financing, but the processes associated with building permits. the inclusion in the comprehensive network contrasts with the social need for this part of the network, according to which the new construction of the high-speed line 27 https://doi.org/10.14311/app.2021.31.0027 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en pavel krupík acta polytechnica ctu proceedings figure 1. ten-t map for rail transport [1]. prague – brno should clearly functionally belong to the core network for passenger transport [4]. high-speed rail lines are also included in the national investment plan, which presents the investment potential of the czech republic until 2050 and which was discussed on 16 december 2019 by the government. it presents a total of seven new track constructions for speeds up to 320 km/h with the date of commencement of implementation from 2025 to 2040 [5]. 2. aims the aim of this article is to describe the possibilities from the perspective of advanced project management methods and the idea of construction 4.0 how not to repeat some significant historical errors in the planning and realization of motorways, which has delayed this process compared to the original plans. 3. methods the article mentions the possible use of advanced project management methods, which the author considers to be one of the important components of construction 4.0. among these advanced project management methods, they work with a predictive project life cycle and different approaches to project management depending on the life cycle. for selected approaches will be discussed the suitability of use for planning and preparation of high-speed rails. 4. results and discussion the pmbok guide [6] states that there is a so-called predictive life cycle, or plan-driven. in this cycle, the scope, time and cost of the project are defined as soon as possible, possible. it is advisable to use it in areas where the final product is known in advance, which is undoubtedly high-speed rails. it also emphasizes the influence and relationship between the project, the project team, and various stakeholders. predictive live cycle of high-speed rail projects in the czech republic shows figure 2. the importance of stakeholders is also indicated by different project management approaches such as iso 21500 standard guidance on project management and project management systems and lean construction philosophy [7]. the situation with the construction of highways only confirms their importance. and it is possible that this will be the case for high-speed rail projects too. the relation between stakeholders and the project in case high-speed rail projects in the czech republic shows figure 3. 28 vol. 31/2021 high-speed rail – highway 2.0? figure 2. predictive live cycle, source: custom. figure 3. the relation between stakeholders and the project, source: custom. 5. conclusion usually it is not possible to name the causes of delays in the construction of motorways constructions. one of culprits is the former minister of transport, who allegedly stopped the preparatory work on dozens of projects in 2010 on the pretext that they are at risk of corruption and because unlimited amounts of money were not available in crisis years. another popular culprit is environmental associations, which continuously delay construction by appealing during any administrative procedure that the construction has to go through before it can even begin. culprit may be also the state bureaucracy as such. for the preparation of buildings it creates such complex regulations that it is beyond its own power to meet their requirements. at the same time, it is unable to reach a civilized agreement with those affected by the construction, from landowners, villages and nature conservation to construction companies. whoever is to blame, the repetition of the highway situation in the case of high-speed rail must be prevented. one of the keys is the consistent identification of the whole predictive life cycle and their stakeholders (including roles). acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs20/100/ohk1/2t/11. references [1] supreme audit office. audit conclusion of audit action 17/05: construction, modernization and reconstruction of motorways. in: supreme audit office: by the autumn of 2017, none of the planned 15 motorway constructions had begun to be built. the blame is the preparation problems [online]. available online at: https://www.nku.cz/cz/pro-media/tiskove-zpravy/ do-podzimu-2017-se-nezacala-stavet-ani-jednaz-planovanych-15-staveb-dalnic-na-vine-jsouproblemy-s-pripravou-id9349/. [2] regulation (eu) no 1315/2013 of the european parliament and of the council of 11 december 2013 on union guidelines for the development of the transeuropean transport network and repealing decision no 661/2010/eu (text with eea relevance). available online at: https://op.europa.eu/cs/publicationdetail/-/publication/58c00e4e-b59c-4b64-ad833d1a02b33fe8/language-en/format-pdfa1a. [3] czech republic. resolution of the government of the czech republic of 22 may 2017, no. 389, on the program for the development of fast railway connections in the czech republic. available online at: https: //apps.odok.cz/attachment/-/down/ihoaamva7kom. [4] ministry of transport of the czech republic. program for the development of fast railway connections in the czech republic. 2017. available online at: https://www.mdcr.cz/dokumenty/strategie/ vysokorychlostni-trate. [5] government of the czech republic. national investment plan of the czech republic 2020-2050. 1. prague: government of the czech republic, 2019. isbn 978-80-7440-244-9. [6] a guide to the project management body of knowledge (pmbok guide). fifth edition. newtown square, pennsylvania: project management institute, [2013]. isbn 978-1935589679 available online at: https://dinus.ac.id/repository/docs/ajar/ pmbokguide_5th_ed.pdf. [7] x. brioso. integrating iso 21500 guidance on project management, lean construction and pmbok. procedia engineering. 2015, 123, 76-84. issn 18777058. available online at: https://linkinghub.elsevier.com/ retrieve/pii/s1877705815031616, https://doi.org/10.1016/j.proeng.2015.10.060. 29 https://www.nku.cz/cz/pro-media/tiskove-zpravy/do-podzimu-2017-se-nezacala-stavet-ani-jedna-z-planovanych-15-staveb-dalnic-na-vine-jsou-problemy-s-pripravou-id9349/ https://www.nku.cz/cz/pro-media/tiskove-zpravy/do-podzimu-2017-se-nezacala-stavet-ani-jedna-z-planovanych-15-staveb-dalnic-na-vine-jsou-problemy-s-pripravou-id9349/ https://www.nku.cz/cz/pro-media/tiskove-zpravy/do-podzimu-2017-se-nezacala-stavet-ani-jedna-z-planovanych-15-staveb-dalnic-na-vine-jsou-problemy-s-pripravou-id9349/ https://www.nku.cz/cz/pro-media/tiskove-zpravy/do-podzimu-2017-se-nezacala-stavet-ani-jedna-z-planovanych-15-staveb-dalnic-na-vine-jsou-problemy-s-pripravou-id9349/ https://op.europa.eu/cs/publication-detail/-/publication/58c00e4e-b59c-4b64-ad83-3d1a02b33fe8/language-en/format-pdfa1a https://op.europa.eu/cs/publication-detail/-/publication/58c00e4e-b59c-4b64-ad83-3d1a02b33fe8/language-en/format-pdfa1a https://op.europa.eu/cs/publication-detail/-/publication/58c00e4e-b59c-4b64-ad83-3d1a02b33fe8/language-en/format-pdfa1a https://apps.odok.cz/attachment/-/down/ihoaamva7kom https://apps.odok.cz/attachment/-/down/ihoaamva7kom https://www.mdcr.cz/dokumenty/strategie/vysokorychlostni-trate https://www.mdcr.cz/dokumenty/strategie/vysokorychlostni-trate https://dinus.ac.id/repository/docs/ajar/pmbokguide_5th_ed.pdf https://dinus.ac.id/repository/docs/ajar/pmbokguide_5th_ed.pdf https://linkinghub.elsevier.com/retrieve/pii/s1877705815031616 https://linkinghub.elsevier.com/retrieve/pii/s1877705815031616 https://doi.org/10.1016/j.proeng.2015.10.060 acta polytechnica ctu proceedings 31:27–29, 2021 1 introduction 2 aims 3 methods 4 results and discussion 5 conclusion acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0131 acta polytechnica 27(0):131–135, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app influencing the indentation curves by the bluntness of the berkovich indenter at the fem modelling jaroslav kovářa, ∗, vladimír fuisa, b, jan tomáštíkc a brno university of technology, institute of solid mechanics, mechatronics and biomechanics, technická 2896/2, 616 69 brno, czech republic b czech academy of sciences, centre of mechatronics, institute of thermomechanics, technická 2896/2, 616 69 brno, czech republic c czech academy of sciences, institute of physics, joint laboratory of optics of palacky university olomouc and institute of physics of the as, na slovance 1999/2, 182 00 praha 8, czech republic ∗ corresponding author: jaroslav.kovar@vut.cz abstract. the influence of the berkovich indenter bluntness on the indentation curves at nanoindentation test of fused silica is determined by the fem (finite element method) in this paper. the berkovich indenter, which is mostly used for the nanoindentation test, was assumed for calculations. the indenter is always blunted in real tests. the bluntness of the indenter has an influence on the results of the nanoindentation test. the combination of the nanoindentation test and the fe modelling can be used for the calculation of the parameters of fused silica plasticity and estimation of the bluntness of the indenter. in this paper, the fem was used for calculating of the indentation curves. two cases of bluntness were assumed. the curves for ideally sharp or blunted indenter were determined. the results of these calculations showed that the indentation curves are dependent on the bluntness of the berkovich indenter. the greatest relative influence of the nanoindentation curves by the bluntness is in the area of small values of the indenter displacement. the impact of assuming blunted indenter to calculated parameters of plasticity was assessed. it is often difficult to measure the bluntness of the indenter but it has an influence on the results of the nanoindentation test and should be included into calculations to make it more precise. keywords: berkovich indenter, bluntness, finite element method, nanoindentation. 1. introduction the nanoindentation test is often used for the testing of materials which cannot be tested by conventional tests. it is often used for the determination of the material parameters of the thin layers or plasticity of the fused silica [1]. the results of the nanoindentation test are often influenced by many conditions. if the fem is used for modelling of the nanoindentation test, these conditions should be included to the calculation to get more precise results. it was proved that the bluntness of the indenter has a significant influence on the results and it should be assessed to get accurate results especially when the displacements of the indenter are similar like height of the blunted part of the indenter. because it is very difficult to measure the real shape of the blunted indenter and use it for fem calculations, it is often modelled like a sphere with known radius but it is unknown if other edges near the tip should be rounded too. in this paper two types of the roundness of the tip of the berkovich indenter were assessed with four radii of the roundness, the fem calculations were done for every case, the results were compared and the best case was chosen. 2. nanoindentation test 2.1. nanoindentation of the fused silica the nanoindentation test can be used for the testing of the material properties of the fused silica. unlike the tension test, in which there is not any significant plastic deformation and the tested specimen fails by brittle fracture after elastic deformation, if the nanoindentation test is used the plasticity occurs in the small volume near the tip of the indenter [1]. fused silica is used in the oliver-pharr analysis as a material for the determination of the shape function of the berkovich indenter [2]. there are few difficulties in the fem modelling, if the fused silica is assumed. firstly, the mechanism of the plasticity of the fused silica is not known yet. the most often used models of stress – strain behaviour are either bilinear [3] or drucker-prager [4]. the bilinear model of the material was chosen for calculations because it should be enough accurate for the main aims of this paper. another problem in the fem modelling of the fused silica nanoindentation was found by the oliver and pharr [5]. the unloading curve should be mostly influenced by the elastic parameters and can be used for the determination of the young’s modulus but if the fem is used the results show steeper unload131 http://dx.doi.org/10.14311/app.2020.27.0131 http://ojs.cvut.cz/ojs/index.php/app j. kovář, v. fuis, j. tomáštík acta polytechnica ing curve than results obtained by the experiment and it leads to the overestimation of the young’s modulus from fem. the reason of this error is still unknown. 2.2. indenter many indenters (like berkovich, spherical, conical etc.) can be used for the nanoindentation [1]. the most used indenter is the berkovich indenter, which was used in this paper. the indenters with the rounded tip like spherical, show steeper loading part of the indentation curve than indenters with the sharp tips (berkovich indenter) [5]. if the bluntness of the berkovich indenter is assumed, the indentation curve is steeper than indentation curve obtained from the calculation with the sharp indenter [6]. this difference is significant in some cases and then the bluntness of the indenter should be included into the calculations otherwise the error in the estimated parameters of the plasticity by the fem could be significant. the two types of the bluntness of the indenter and few radii of the bluntness of the berkovich indenter will be assessed in this paper and the indentation curves from the fem will be compared with the indentation curve obtained by the experiment. 3. methods the fem was used for the determination of indentation curves. 3.1. models of geometry the berkovich indenter pressing into the specimen from the fused silica was assumed for calculations. the size of the indenter and the specimen was chosen to be much bigger than the displacements of the indenter (fig. 1). then the influence of the edge of the specimen on the results is small. only one sixth of the specimen and the indenter was modelled because there are three planes of symmetry. then the bluntness of the indenter was modelled. two types of the bluntness were assumed. in the case a (fig. 2), the tip of the indenter and the edges of the indenter were blunted. in the case b (fig. 3) another roundness, with the same radius as in the case a, was applied on the point where intersect the lines at the border of the roundness. for each case of bluntness, four radii of the bluntness were assessed. the values of the roundness were 100 nm, 200 nm, 400 nm and 800 nm. to compare the results, the case without bluntness was calculated too. 3.2. boundary conditions the indenter was loaded by the displacement of the specimen in two loadsteps. in the first loadstep, the displacement 612 nm plus the gap between the specimen and the blunted indenter against the way of the z axis was applied at the upper area of the indenter. in the second loadstep, the unloading by the displacement of the indenter back to the 200 nm was done. the displacement of the lower area of the specimen in figure 1. geometry of the indenter and specimen [mm]. figure 2. mesh of the tip of the indenter with the bluntness 200 nm, (case a, red line shows the border of the blunted area, yellow the border between the blunted tip and the edge). figure 3. mesh of the tip of the indenter with the bluntness 200 nm (case b, red line shows the border of the blunted area, yellow the border between the blunted tip and the edge). the way of the z axis was fixed. on the areas of the symmetry, the displacements perpendicular to these areas were fixed. 132 vol. 27/2020 fem modelling of indentation curves 3.3. models of material the diamond indenter was modelled with the linear elastic model of material with the young’s modulus e = 1 141 000 mpa and the poisson’s ratio µ = 0.07. the fused silica was modelled with the bilinear model of material. the elastic parameters were taken from [1] (e = 72 000 mpa, µ = 0.17). the parameters of plasticity (yield strength and hardening) were estimated from the experimental data by this procedure. from the comparison of the experimental data and the results from the fem, the bluntness of the indentor in the case b was determined like 200 nm. the indentation fits well in the area of the low loads where the plasticity is not so important. then the yield stress was estimated like re = 4000 mpa. the hardening of the material was not assumed. 3.4. mesh and the fem calculation the mesh was made coarse at the edge of the specimen but very fine at the centre to obtain precise results (fig. 4). similar procedure was used on the indenter (fig. 2 and 3). the elements of the type solid 185 were used on the specimen and the quadratic solid 186 elements on the indenter to describe the edge of the indenter more precise. between the indenter and the specimen, the contact region was made and conta 174 / targe 170 elements were used. to obtain more accurate results, the penetration tolerance factor was reduced to 0.01. both loadsteps were divided into a few substeps and the large displacements algorithm was applied. because of the small values of the displacements, the force convergence criterion was lower to 0.00001, to obtain more precise results. the calculation with this algorithm was done for every value of the bluntness of the indenter. figure 4. the mesh in the area near contact. 4. results and discussion the indentation curves (fig. 5) were obtained as the results of the fem calculations. the curve obtained by the model without shape imperfection showed the least steep loading curve and unloading curve too. the steepness of the curves grew with the bluntness. this agreed with the oliver-pharr analysis, where the higher steepness of the unloading curve is taken into account by the indenter shape function. if the sharp indenter was assumed, the unloading curve was still steeper than the experimental. this error was firstly observed by [5] and it is not explained yet. it could be caused by not enough accurate model of material but this was not important for this paper, which was focused on the assessing of the influence of the bluntness of the indenter. the difference between the results with the bluntness a (blunted only the tip and the edges fig. 2 and the bluntness b (blunted the tip, the edges, and rounded points, fig. 3) was small, if the bluntness was low but it grew with the growing bluntness. the relative differences of the maximal values of the force related to the bluntness were: 1.54 % for bluntness 100 nm, 2.92 % for 200 nm, 5.78 % for 400 nm and 11.4 % for 800 nm. the bluntness b should be nearer to the reality because the edges should be blunted at least a little but for new indenters, which are not so much blunted, the difference between these models was smaller than 2 %. when the bluntness b with radius 200 nm was assumed, the loading curve obtained from the fem calculation fitted the curve from the experiment. the similar but a little less accurate results were obtained by the bluntness a with radius of 400 nm. the bluntness influenced every part of the curve because it shifted the curve except the lower part of the curve which was mostly influenced because there is the spherical part of the tip of the indenter (fig. 6). the steepness of the curve in this area dramatically grew with the growing bluntness of the indenter. this fact is in the agreement with the results of the indentation curves obtained by the sharp indenters like berkovich and spherical indenter. the indentation curves obtained by the spherical indenters show steeper indentation curves [6]. from the fig. 6 is clear that the bluntness b with radius of the 200 nm was very near to the experimental curve even in the area of the low displacements. the similar relative error showed bluntness a with radius 400 nm. other cases show bigger relative errors in the area of the small displacements than in the maximal values of the forces. because the influence of the plasticity is smaller in this area than in the end of loading and the influence of the bluntness is higher, the agreement of the curves from the experiment and the fem calculation in the area of the small displacements is great criterion for the determination of the bluntness of the indenter for the next calculations. 133 j. kovář, v. fuis, j. tomáštík acta polytechnica figure 5. indentation curves for all cases. figure 6. indentation curves in the area of small displacements (detail of the fig. 5). 5. conclusions the two types of the bluntness of the berkovich indenter with four radii of the roundness have been compared in this paper. the loading and the unloading curves were getting steeper while the radius of the roundness grew. the bluntness b, in which are the edges of the indenter, the tip of the indenter, and the points, where the edges intersect, blunted, gave steeper indentation curves than bluntness a, in which are blunted only the edges and the tip of the indenter. the indentation curve for the indenter with the bluntness b with radius 200 nm was the closest to the experimental curve. there still remained the problem with higher steepness of the unloading curve in the results. it could be caused by the inaccuracy of the used model of material. the bluntness had the biggest impact in the area with the small displacements, where the blunted spherical tip caused higher steepness of the indentation curve. elastic parameters are known and the fitted parameters of the plasticity have only small impact in the area of the small displacements. then this could be used as a great criterion for the determination of the bluntness of the indenter with the fem in the future work. acknowledgements the experimental data were measured by the institute of physics of the czech academy of sciences, joint laboratory of optics of palacky university olomouc and institute of physics as cr. 134 vol. 27/2020 fem modelling of indentation curves references [1] j. kovář, v. fuis. fem simulation of the nanoindentation test with rigid and non-rigid indenter. engineering mechanics 2019 pp. 189–192, 2019. doi:10.21495/71-0-189. [2] w. oliver, g. pharr. an improved technique for determining hardness and elastic modulus using load and displacement sensing indentation experiments. journal of materials research 7(6):1564–1583, 1992. doi:10.1557/jmr.1992.1564. [3] d. torres-torres, j. muñoz-saldaña, l. a. g.-l. de guevara, et al. geometry and bluntness tip effects on elastic-plastic behaviour during nanoindentation of fused silica: experimental and fe simulation. modelling and simulation in materials science and engineering 18(7):075006, 2010. doi:10.1088/0965-0393/18/7/075006. [4] k. gadelrab, f. bonilla, m. chiesa. densification modeling of fused silica under nanoindentation. journal of non-crystalline solids 358(2):392–398, 2012. doi:10.1016/j.jnoncrysol.2011.10.011. [5] s. shim, w. oliver, g. pharr. a comparison of 3d finite element simulations for berkovich and conical indentation of fused silica. international journal of surface science and engineering 1(2/3):259, 2007. doi:10.1504/ijsurfse.2007.015028. [6] n. yu, a. a. polycarpou, t. f. conry. tip-radius effect in finite element modeling of sub-50 nm shallow nanoindentation. thin solid films 450(2):295–303, 2004. doi:10.1016/j.tsf.2003.10.033. 135 https://doi.org/10.21495/71-0-189 https://doi.org/10.1557/jmr.1992.1564 https://doi.org/10.1088/0965-0393/18/7/075006 https://doi.org/10.1016/j.jnoncrysol.2011.10.011 https://doi.org/10.1504/ijsurfse.2007.015028 https://doi.org/10.1016/j.tsf.2003.10.033 acta polytechnica 27(0):131–135, 2020 1 introduction 2 nanoindentation test 2.1 nanoindentation of the fused silica 2.2 indenter 3 methods 3.1 models of geometry 3.2 boundary conditions 3.3 models of material 3.4 mesh and the fem calculation 4 results and discussion 5 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0638 acta polytechnica ctu proceedings 33:638–643, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague rational and sustainable procedure in the design of bridges using pre-stressed concrete beams a theoretical practical method abel noé xochicale cortés bufer & co. méxico, sa de cv., privada 9 b sur #5117, prados agua azul, 72430 puebla, pue., mexico correspondence: anxochicale@yahoo.com abstract. currently there is not a direct procedure for the design of this bridge system, the procedure is still based on trial and error, there are no studies that give us both the optimum number of beams and other characteristics such as the correct thickness of the slab or its adequate strength. keeping in mind the concept of optimization, which supports the idea of sustainability and environmentally friendly structures, we apply this concept to the design of pre-stressed concrete girder bridges, which are an integral part of the road network in most countries of the world. knowing also that concrete is the most produced material by man and that without it, it would be difficult to understand our current world, due to its own characteristics that make it unique when it comes to sustainability, however, its production requires energy and material consumption, that is why an optimal design of a superstructure will lead to the minimum consumption of materials obtaining a more friendly design with nature. it is a global necessity to obtain and deliver concrete structures that meet the conditions of sustainability. keywords: bridges, girders, sustainability. 1. introduction it is true that after many years in which have been used pre-stressed concrete girders supporting a concrete slab, has not been able to establish the suitable type of girder according to the span, nor to know the maximum performance of each type of girder, giving in this way many solutions to a given problem, but none adopted in a generalized way to provide an appropriate and economical solution to the proposed array of girders, in other words, the most sustainable or optimum solution. it is shown a direct method to determine the number of girders more suitable according to the type to use. 2. review to the technical literature 2.1. relationship between optimization and sustainability recently has emerged the concept of the sustainable design of structures, which by applying it to prestressed concrete girder bridges, could be synthesized in the use of 1. minimum number of girders, 2. optimum thickness of slab over girders, 3. minimum number of columns in the case of piers 4. minimum number of spans in the bridge, which is achieved using the girders to its maximum performance. in 1971, f. jacques, [1] conducted a study motivated by the need to have greater spans with the existing girder sections and to have the most suitable section for local conditions. the study gave the correct importance to the employment of segmented girders spliced in field with the application of a posttensioning, however due to the little studies on this process is opted to find another solution. although we had enough information about the types of girder sections in use, the question remained is, what it is the cheapest and practical section to be used?. it was considered the following criteria; • practicability, realistic concrete strengths 6000 psi (42 mpa) consistently, perhaps up to 7000 psi (50 mpa), realistic size and weight limits for precast girders. • safety, the use of a single unit precast section appears desirable since it can span existing roadway without shoring. this eliminates inherent shoring hazards. • esthetics, it appears desirable to eliminate the stubby end blocks such as on the old aasho sections. also, it seems desirable to eliminate the possibility of texture blemishes that can result from field splices. • economy, the cost of field splicing, in all cases, appears to be an added cost that can be justified only if a single piece section cannot be hauled overroad to the site. in terms of number of girders and their spacing, it is clear that a deep girder is capable of carrying a greater payload moment when compare to a shallow girder for a given span, this allows that greater 638 https://doi.org/10.14311/app.2022.33.0638 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 bridges design using pre-stressed concrete beams figure 1. optimum girder spacing cost curves for seven stone concrete girders. height beams are spaced at greater distance than the shallower girders which results in fewer girders for a width of bridge, but this is reflected in a thicker deck slab with more reinforcement, on the other hand, may require greater volume of earthwork for the clearance require, several sections and parameters were defined like, span and girder spacing, as well as, the thickness of the slab deck. two concrete strength in girders were defined, 6000 y 7000 psi. live load aasho hs20-44, intermediate diaphragms at 12 m maximum and 20 cm thick have been proposed and using a computer program, were obtained graphs that relate the span, type of girder and cost per unit of surface, as we see in the figure 1. as a first conclusion of this study we can say that for a certain span, optimum spacing between girders provides more economic bridge, in this case, a spacing equal to 1.25 times the height of the girder. it is noted the concept of balanced section according to the area of lower flange compared to the top flange area. 2.2. advances in cross section optimization in the report by rabbat and russell [2], they wanted to determine optimal designs for common girders and to analyze the potentialities of standardizing these sections as well as make recommendations for practical and economic designs. to do this we studied structural efficiency and cost-effectiveness of the best existing designs and study the impact to make some modifications to aashto girders, it was proposed to study the following sections shown in the figure 2 making the following assumptions • design conforms to aashto specifications. • live loads consists of hs 20-44 loading. • girders are simply supported • a typical interior girder is considered concrete deck is cast in place and acts compositely with the girder. deck formwork is supported on the girder. in calculations of the composite section properties, the transformed area of strands is neglected. concrete compressive strength of the deck is constant and equal to 5000 psi (28 mpa) at 28 days. strands are grade 270 (1,900 mpa) stress relieved with 1/2 in ( 1.3 cm) diameter and have an idealized trilinear stress-strain curve. total pre-stressing looses are constant and equal to 45,000 psi (310 mpa). initial or long term camber or sag do not govern design, because aashto specifications do not specify deflections limits for concrete bridges. as a result of studies in different sections of girders show that the aashto sections are not the most economical for the characteristics of materials, the live load applied and to the design conditions prevailing in the country, and suggests to modify the above sections to reduce cost in the work, also suggests the use of concrete in beams from 35 to 50 mpa to increase the until span about 15%. also suggests that the modified section "bulb-t" is most suitable for spans from 24 to 42 meters, for greater spans it is recommended spliced sections. intermediate diaphragms are not needed, end diaphragms are sufficient. 2.3. maximum structural performance of the cpci girder in 1993, lounis and cohn [3], defined first some of the parameters involved in the design of a bridge with girders, such as: a) live load, b)the type of girder, c) the strength of the concrete slab and the strength of concrete in the girder and 639 abel noé xochicale cortés acta polytechnica ctu proceedings figure 2. modified girder analyzed (aashto, bulb t, colorado and washington). d)existence or not of continuity at intermediate supports and with the use of a computer program defined for spans and common widths, the type of girder, the minimum number for girders and the maximum spacing for girders, maximizing the performance of these elements. 3. a practical-theoretical approach based on the study of the structures previously shown and using basic concepts of optimization, has been related numerically the different parameters involved in the design of this bridge system, defining certain parameters as is mentioned in previous reports, has been obtained an optimization constant involved in the geometry of the superstructure, this constant is an indicator which depends of the type and number of girders, so that, thickness and concrete strength of the slab over girders. at the same time that this constant produces an optimal superstructure, implicitly define the more sustainable superstructure. it has been established the following equation: η × k0 × h0! 0 dy l × b → {k1 → 0.02} (1) where • η . . . represents he number of girders in the superstructure. • k0 . . . is a constant with a value 1.0 approximately for girders with web width of 20 cm. • l . . . represents the span. • b . . . is the width of the superstructure. • k1 . . . is the optimization constant with a value of 0.02 approximately. • h0! 0 dy . . . represents the girder depth in the superstructure. it has been established certain constraints; • concrete strength in girders of f ′c = 40, 45 and 50 mpa • two concrete strength are considered for the slab 25 and 30 mpa 640 vol. 33/2022 bridges design using pre-stressed concrete beams figure 3. maximum feasible girder spacing, optimum girder spacing and optimum number of girders for single, two and three-span continuous cpci girder bridges. z. lounis and m. cohn. • live loading is that defined by sct t3-s2-r4 (710 kn), which is very similar to the o.h.b.d. live loading. • design specifications is that one given by aashto. • width web is approximately 20 cm. this is summarized in the figure 4, which relates the live loads and the type of girder, obtaining the optimal ratio of the sum of the height of beams on the system, divided by the product of the span and total width of the system. the kxoc factor should be close to the value of 0.02. any system which produce a kxocfactor close to this value, will produce an optimal solution. a very remarkable feature of this factor, is that it is bidirectional, because it can be applied to both directions of the axis of the bearings. design process: • according to the live load and the girder type selected, choose the appropiate constant kxoc. • get the product of the constant kxoc times the span and the width of the superstructure. • choose the more appropriate girder according to the its depth htr . • obtain the number of girders dividing the first product by girder depth, closing it to the next higher integer, obtaining this way, the real value for kxoc. • accommodating the number of girders obtained above, we get the spacing between girders which must be within the acceptable maximum values. • the final design of the girder will be according to the initial conditions. applying the formulation for the following cases, we obtain a-factor kxoc equal or very close to 0.02, so reaching the optimal solution. this formulation perfectly agrees with the results given in figure 4 for the condition of single spans girder bridges. for continuous spans we can take into account only 0.85 of the length between intermediate supports. for the case of the reconstruction of the walnut lane bridge, we can notice that we have a kxoc = 0.0202 (see [4]). it is necessary to mention another determining factor in the cross section of the superstructure called kbr , which can be defined as the result of dividing the sum of the depth of the beams in the superstructure by the width of the same superstructure, as explained below (see xochicale et al. [5]): 0.4 < η × k0 × h0! 0 dy l × b < 0.8 (2) 4. application of the concept of sustainability to this bridge type in summary, a sustainable design is one that is obtained at this time without compromising the ability of future generations to solve their own needs. from this definition, an engineering and sustainable project such as a bridge project based on concrete beams, is one that is conceived, designed, built, operated on, is maintained and eventually is put out of service so that these activities require as little as possible in terms of energy and material consumption in support of the community. under this concept of sustainability, what is sought are engineering projects 641 abel noé xochicale cortés acta polytechnica ctu proceedings figure 4. graphic definition of kxoc, according to live loading and girder type. figure 5. definition of sustainability triple bottom line. that have the greatest positive impact at the intersection between the interests of the people they serve, the planet in its short and long term scale and the utility for anyone involved in the life of that project as defined by the concept of triple bottom line. a fully sustainable bridge is one that strives to serve the population for which it was made, and that its long-term cost is totally environmentally friendly. in this sense, concrete plays an important role in providing solutions to future challenges by contributing to the construction of sustainable infrastructure. with a growing population, concrete’s crucial contribution is required as it supports communities in building their infrastructure to connect cities and transport people and goods. its benefits and sustainable characteristics make concrete a unique material for a more environmentally friendly future. the above according to n. subramanian [6]. 5. application to a current project in mexico applying these concepts to a current project to be built in mexico, in the state of michoacán, we can see the advantage of applying the concepts of optimization and sustainability to this type of structure. this is a road distributor which consists of a series of 13 spans of 34.0 meters, in which 80 aashto type v beams are used, and 19 spans of 29.0 meters, in which 99 aashto type iv beams are used, meaning the following report in terms of material consumption, each span contains 5 beams spaced at 150 cm. and is reduced to 4 beams spaced at 200 cm. in its modified semi-optimized version in which it has a kxoc closer to 0.02. the final version is the optimized one and shows the lowest consumption of materials whose kxoc factor is even lower than 0.02. it is noted that with the application of the optimization criterion it is possible to reduce quantities of materials by 20% and with it the total costs of the work, having this way a structure that fulfills the aspects of sustainability. 6. materials savings in general terms it can be seen that for the mentioned project it is possible to save up to 980 m3 of concrete and up to 214 tons of steel, this represents to avoid up to 340 tons of co2 in the production of the cement, knowing in addition that at the moment up to 75% of the steel is recycled in the construction, it is immediately noticeable that an optimal design leads to a sustainable design since it means savings in materials and the reduction of pollutants, in addition to 642 vol. 33/2022 bridges design using pre-stressed concrete beams figure 6. study case to show the kxoc factor and the relation between the number of girders and the consumption of materials. other expenses as they are the transfer of equipment, personnel and other materials. 7. conclusions the shown method has given very successful solutions and it is recommended in the development of new projects. once defined the type and the number of beams by use, spacing is defined at the same time, respective calculations done can proved the convenience of using each type of girder that fits its height to the closest value to the factor kxoc = 0.02. in addition, as can also be seen, this value leads to economic and rationally designed solutions including sustainability features, among which can be mentioned, the reduction in the consumption of natural resources, the reduction of pollutants emitted into the environment and the possibility of recycling the materials used. references [1] f. j. jacques. study of long span prestressed concrete bridge girders. pci journal 16(2):24-42, 1971. https://doi.org/10.15554/pcij.03011971.24.42. [2] b. g. rabbat, h. g. russell. optimized sections for precast prestressed bridge girders. pci journal 27(4):88-106, 1982. https://doi.org/10.15554/pcij.07011982.88.106. [3] z. lounis, m. z. cohn. optimization of precast prestressed concrete bridge girder systems. pci journal 38(4):60-78, 1993. https://doi.org/10.15554/pcij.07011993.60.78. [4] c. c. zollman, f. depman, j. nagle, et al. a 40-year saga building and rebuilding of philadelphia’s walnut lane memorial bridge. pci journal 37(4):66-82, 1992. https://doi.org/10.15554/pcij.07011992.66.82. [5] a. n. xochicale, reg. no. 03-2017-021011304600-1 optimizacion de super-estructuras y aplicacion de trabes de sección i de concreto presforzado en puentes carreteros de grandes claros para sistemas de losa sobre trabes, 2017. [6] n. subramanian. sustainability challenges and solutions. the indian concrete journal, 2007. https://www.researchgate.net/profile/subramani an-narayanan-2/publication/290228957_sustainabi lity_-_challenges_and_solutions/links/61b4f e5463bbd9324286718a/sustainability-challenges-a nd-solutions.pdf. 643 https://doi.org/10.15554/pcij.03011971.24.42 https://doi.org/10.15554/pcij.07011982.88.106 https://doi.org/10.15554/pcij.07011993.60.78 https://doi.org/10.15554/pcij.07011992.66.82 https://www.researchgate.net/profile/subramanian-narayanan-2/publication/290228957_sustainability_-_challenges_and_solutions/links/61b4fe5463bbd9324286718a/sustainability-challenges-and-solutions.pdf https://doi.org/10.14311/app.2022.33.0398 acta polytechnica ctu proceedings 33:398–403, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague interfacial bond behavior of adhesively-bonded timber/cast in situ concrete (wet bond process) ali nemati giva, qiuni fua, libo yana, b, ∗, bohumil kasala, b a technische universität braunschweig, department of organic and wood-based construction materials, hopfengarten 20, 38102 braunschweig, germany b fraunhofer wilhelmklauditz-institut, center for light and environmentally friendly structures, bienroder weg 54e, 38108 braunschweig, germany ∗ corresponding author: l.yan@tu-braunschweig.de, libo.yan@wki.fraunhofer.de abstract. the goal of this research was to study the strength of the interfacial bond between cast-in-situ concrete and engineered timber (cross-laminated timber (clt)). double lap specimens were manufactured using fresh concrete that was cast between two clt blocks. polyurethane and epoxy adhesives were used to bond the wet concrete with the clt blocks. the shear strength of wet-bond specimens was compared with the specimens prepared under dry conditions (prefabricated concrete cube glued to clt blocks). the statistical analysis (t-test) of bond strength showed that the shear strengths of wetand dry-bond specimens using epoxy and polyutrthane adhesives were no significantly different for the tested c25 plain concrete and the clt. the failure mode of dry-bond specimens were concrete failure near the interface, however, debonding at interface was the dominant failure for the wet-bond specimens. keywords: cast-in-situ concrete, dry bond, shear strength, timber-concrete composite structures, wet bond. 1. introduction timber-concrete composite system is a construction technique to strengthen and stiffen the existing timber floors and new construction parts such as decks in short-span bridges and multi-story buildings. this technique requires the connection system to transfer shear stress between timber and the concrete composite system. rigid and strong connection system maximize the composite action. rigidity of connection system will increase when notches or holes [1] are cut/drilled in the timber or a continuous connector like adhesive is used [2]. application of adhesive has recently been studied in bonding timber and concrete due to its significant advantages including: a) uniform stress distribution over the bond area, b) removal of cutting and drilling in the wood substrate and c) reduction of workmanship and cost [3]. satisfactory stiffness and strength is of importance for timber-concrete structures. besides the stiffness and strength, other aspects should be considered such as quality, simplicity and speed of manufacturing and erection. two approaches can be taken to manufacture adhesively bonded timber-concrete elements: • adhesively bonded timber-prefabricated concrete composite (dry bond). • adhesively bonded timber-cast-in-situ concrete composite (wet bond). in a dry bond, adhesively bonded timber-concrete beam/slab is fabricated in a workshop under the controlled environment, tranported to cosnstrution site, lifted up and placed on a desired location. in case of wet bond, the wet concrete is directly pumped to wood surface covered with glue. the wet bonding has its own merits such as: • easy transport of wet concrete to desired location making time and cost savings especially for multistory buildings (removal of crane application). • elimination of gap and discontinuities in the adhesive interface area. there is a concern regarding the insufficient bond strength between timber and cast-in-situ concrete when it is compared with dry bond. when the fresh concrete is poured on the wet adhesive, it can cause the adhesive movement, incomplete adhesive curing and consequently bond strength reduction [8, 9]. few researchers studied the bond behaviour of timbercast-in-situ concrete connected with glue. brunner et al [9] investigated the wet bond behaviour of glue laminated timber (gl24h) and cast-in-situ concrete (c25/30) glued with epoxy adhesive. they concluded that the production of wet bond in timber-concrete composite structure is delicate due to danger of adhesive movement during pouring fresh concrete. clouston et al [10] experimentally studied the influence of epoxy adhesive on strength and stiffness improvement of full-scale wood plank-cast-in-situ concrete composite floor decks. the strength and stiffness were increased by about 4 and 2.2 times compared to woodcast-in-situ concrete composite with no connection 398 https://doi.org/10.14311/app.2022.33.0398 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 adhesively-bonded timber & cast in situ concrete clt bending strength (mpa) 18 modulus of elasticity in fiber direction (gpa) 12 shear modulus (mpa) 690 density (kg/m3) 527 table 1. mechanical and physical properties of clt [4, 5]. mass percentage cement water fine aggregates(sand) medium aggregates coarse aggregates normal concrete 1 0.6 2 1.27 2.73 aggregates size (mm) − − 0 − 2 2 − 8 8 − 16 table 2. designed mixing ratio for concrete. figure 1. concrete mixing process. epoxy [6] polyurethane [7] commercial name sikadur 300 semparoc 12 nv density 1.16 g/cm3 1.25 g/cm3 viscosity 700 mpa.s 8000 − 12000 mpa.s pot life (at 23◦ c) 4 hours 15 minutes curing time 7 days 7 days table 3. basic information of epoxy and polyurethane adhesives used in this study. in the interfacial area. additionally, glue as a continuous connector in the interface area changed the composite failure mode from concrete cracking and debonding between wood and concrete components to interfacial shear failure by sliding concrete over the wood planks. this paper compares the shear bond strength of two possible manufacturing techniques of adhesivelybonded timber-concrete composite namely wet and dry bonds. two wood blocks were glued with the prefabricated and cast-in-situ concrete cubes using epoxy and polyurethane adhesives. after adhesive and concrete curing, static shear tests up to failure were carried out up to failure. fracture surfaces of failed specimens were analysed to investigate possible correlation with shear bond strength. 2. experimental work 2.1. material to fabricate timber-concrete glued joints, cross laminated timber (see table 1) with a commercial product name of bbs 125 3-s was used. the clt wood was made of norway spruce and was composed of three layers with thicknesses of 20,40 and 20 mm, respectively (totally 80 mm). prefabricated and cast-in-situ concrete have the same mixing ratios (see table 2) and were fabricated using the same mixer. portland cement cem i 42.5 n was used for the concrete mixing. in production of normal concrete, the water was added into two steps (see. figure 1) to ensure aggregates were wet enough causing stronger interfacial transition zone between concrete aggregates and mortar matrix. one-component polyurethane and two component epoxy adhesives applicable to wood and concrete substrates were used in this study. some parameters of these products are shown in the table 3. 2.2. double lap joint geometry double lap joint is symmetrical about the mid plane of a specimen, therefore, bond rotation and the amount of peel stress is considerably reduced compared to equivalent single lap joint [11]. figure 2 shows the designed double lap joint with dimensions and material location. 2.3. double lap joint preparation 2.3.1. dry bond preparation for the dry bond fabrication (see figure 3), fresh concrete was poured into mold, vibrated and cured for 24 hours. afterwards, the concrete blocks were immersed in the water for 6 days at temperature of 23◦ c. finally, they were taken away from water tank and placed in the climate chamber (65% 399 a. nemati giv, q. fu, l. yan, b. kasal acta polytechnica ctu proceedings figure 2. double lap joint geometry. pouring fresh concrete in the mold placement of concrete blocks in the water tank storing concrete blocks in the climate chamber after 1 day after 6 days after 21 days clt clt gluing first concrete block to first wood substrate gluing wood-concrete joint to second wood substrate adhesive layer concrete wood-concrete joint after adhesive consolidation after adhesive consolidation timber-concrete glued joint clt clt clt figure 3. procedure of dry bond preparation. rh and 23◦ c) for completion of concrete curing (for the rest of 21 days) based on the german concrete standard (din en 12390-2 [12]). then, the concrete and clt wood blocks were bonded together and cured for seven days according to adhesives technical datasheets [6, 7]. 2.3.2. wet bond preparation in the wet bond fabrication, the wood blocks were placed next to each other separated by plywood plates. the surfaces of wood blocks were covered by epoxy and polyurethane adhesives (see figure 4a), turned around, fixed by horizontal and vertical c-clamps on their positions and fresh concrete was poured in the middle of the wood blocks (see. figure 4-b). timber-concrete double lap joints were released from mold after 24 hours and then the concrete cubes were wrapped up by damp fabrics to avoid incomplete concrete curing process. finally, the specimens were stored in the climate chamber (65% rh and 23◦ c) for the rest of 27 days. 2.4. shear test setup the shear test of wet bond and dry bond specimens were carried out using din-en 392 standard [13] proposed for timber glued joints. wood blocks were placed onto the steel plates and constrained from rotation and horizontal movement using two steel plates and c-clamps (see. figure 5). the load was forcecontrolled with a loading rate of 30 kn/min through the steel plate positioned on the concrete. the double lap joint specimens were tested up to failure and shear strength of wet and dry bonds was calculated according to eq. 1: 400 vol. 33/2022 adhesively-bonded timber & cast in situ concrete figure 4. manufacturing process of wet bond a) pouring and spreading glue on the wood surface b) pouring fresh concrete between wood blocks covered with glue. figure 5. shear test setup for measurement of shear strength of wet and dry bonds. τu = fu 2a (1) in which, τu, f and a are shear strength, ultimate shear load and adhesive bond area, respectively. τu thus represents the average shear strength for both faces of the concrete block. 3. shear test results the average values and standard deviations of wet and dry bond specimens are shown in table 4. the shear strengths of each adhesive type in wet and dry bond were compared using t-test (two sample assuming unequal variances) to reveal whether their expected mean values (µ) were equal (h0 = µwet bond = µdry bond). the two-tail p-values of wet and dry bonds for epoxy and polyurethane adhesives are shown in table 5. this value indicates the probability of rejecting or accepting the zero hypothesis (equal mean value for wet and dry bonds). according to table 5, the two tail p-values of epoxy (p = 0.88) and polyurethane (p = 0.071) adhesives are higher than the level of significance (α = 0.05) indicating no significant difference between mean values of wet and dry bonds by acceptance of zero hypothesis. it should be pointed out that the significance level, α, is set somewhat arbitrarily but the value of 0.05 represents a reasonable probability of the error in conclusions. as it is evident from the table 4 that the epoxy adhesive had higher shear strength compared to polyurethane adhesive in wet and dry bonds so that the shear strength increased about 98.2% and 58% in wet and dry bonds, respectively. it is worth mentioning that the increase of shear strength achieved by epoxy adhesive does not diminish the application of polyurethane adhesive. in real application when timber-concrete composite beam/or deck is subjected to flexural loading, 401 a. nemati giv, q. fu, l. yan, b. kasal acta polytechnica ctu proceedings type of joint epoxy polyurethane dry bond τu (mpa) 2.26 1.43 sd (mpa) 0.72 0.43 number of specimens 8 11 wet bond τu (mpa) 2.22 1.12 sd (mpa) 0.49 0.27 number of specimens 8 8 table 4. shear strength (τu) and standard deviation (sd) of timber-prefabricated and timber-cast-in-situ glued joints. type of joint epoxy polyurethane two tail p-value 0.88 0.071 level of significance (α) 0.05 0.05 reject or accept the zero hypothesis (h0) accept accept table 5. two-tail p-value results from wet bond versus dry bond. wet bonddry bond polyurethane epoxy 50 mm 50 mm 50 mm 50 mm figure 6. fracture surfaces of timber-prefabricated concrete and timber-cast-in-situ concrete glued with polyurethane and epoxy adhesives. polyurethane adhesive as a ductile adhesive (i.e. adhesive can undergo plastic deformation before failure) can well perform against adhesive peel stress, which may cause premature failure in the adhesive bondline. furthermore, polyurethane is cost saving and compatible with wood substrates [14–16], therefore, there is a motivation for further investigation on the polyurethane adhesive in wet and dry fabrications. 3.1. fracture surface of wet and dry bond specimens the fracture surfaces of failed specimens of wet and dry bonds are illustrated in figure 6. as it is evident from figure 6 that both polyurethane and epoxy adhesives, the amount of concrete failure remained on the wood blocks was significant in dry bond specimens. however, debonding at interface was the dominant failure mode in the wet bond specimens. the failure mode change from concrete failure to debonding at interface is a clear indication of weaker shear strength in wet bond application [17–19]. 4. conclusion in this paper, the influence of two main types of concrete fabrication including cast-on and -off sites on shear strength of timber-concrete glued joints was studied. two types of adhesive polyurethane and epoxy were used in this work. in timber-prefabricated concrete glued joint as a dry bond, the concrete was precast and adhered to wood blocks by its own weight while for timber-cast-in-concrete glued joint (as a wet bond), fresh concrete was added between two wood blocks covered with wet adhesive. the results showed that the shear strength of clt-prefabricated concrete and clt-cast-in-situ concrete bonded with epoxy adhesives had no significant difference whereas for case of polyurethane adhesive, the wet bond shear strength reduced by about 21.6% compared to dry bond. in addition, fracture surfaces of wet and dry bond specimens for epoxy and polyurethane adhesives were investigated. the change of failure mode was observed from concrete failure in dry bond to debonding at interface in wet bond which was an evidence in reduction of wet bond shear strength. 402 vol. 33/2022 adhesively-bonded timber & cast in situ concrete in manufacturing wet bond (especially timberconcrete composite deck/beam), it is recommended that the fresh concrete is poured at several locations of timber beam. this is due to distribute adhesive amount along the timber beam/deck. brunner et al [8] suggested that the height of pumping fresh concrete would be low enough to avoid the adhesive movement. increasing the height of pumping/or pouring fresh concrete may cause to adhesive splashing and movement in the location of pumping/or pouring fresh concrete. therefore, no adhesive remains for gluing timber and concrete. moreover, gentle use of vibrator is approved for debubbling fresh concrete, however, vibrator should not have any contact with adhesive layer. acknowledgements the research was financially supported by fachagentur nachwachsende rohstoffe e. v. (fnr, agency for renewable resources) founded by bundesministerium für ernährung und landwirtschaft (bmel), under the grant award no.: 22011617, and by bundesministerium für bildung und forschung (bmbf) (grant no.: 031b0914a). references [1] a. heiduschke, and b. kasal, composite cross sections with high performance fiber reinforced concrete and timber, forest products journal 53(10):74-78, 2003. [2] d. e. c. yeoh, behaviour and design of timber-concrete composite floor system, university of canterbury. department of civil and natural resources, ph.d. thesis, 2010. [3] t. tannert, b. endacott, m. brunner, et al. long-term performance of adhesively bonded timber-concrete composites. international journal of adhesion and adhesives 72:51-61, 2017. https://doi.org/10.1016/j.ijadhadh.2016.10.005. [4] binderholz-gmbh. binderholz brettsperrholz bbs, 2017. https: //www.binderholz.com/fileadmin/pdf/services_ko ntakt/videos_download/prospekte/bbs_d_web.pdf. 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https://doi.org/10.14311/app.2022.33.0417 acta polytechnica ctu proceedings 33:417–423, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague glued timber concrete composite walls using ultra-high-performance concrete thomas oberndorfera, ∗, frank hungerb, oliver fischera a technical university of munich, department of civil, geo and environmental engineering, chair of concrete structures, theresienstraße 90, 80333 munich, germany b technical university of munich, department of civil, geo and environmental engineering, chair of timber structures and building construction, arcisstraße 21, 80333 munich, germany ∗ corresponding author: thomas.oberndorfer@tum.de abstract. this paper gives a short introduction of the development of cross laminated timber (clt) concrete composite walls with a glued connection. in the composite walls, prefabricated ultra-highperformance concrete (uhpc) lamellas replace timber lamellas at the core layer of the clt elements. to check the feasibility of gluing timber to uhpc small-scale shear, delamination and bonding tests were performed and showed promising results. the load bearing behaviour was analysed with centrically and eccentrically loaded tests on wall segments. analytic modelling of the wall experiments using an effective bending stiffness, based on shear analogy method, showed a good correlation to the experimental results. keywords: adhesive bond, cross laminated timber, timber concrete composite, ultra-highperformance concrete. 1. introduction cities are forced to grow, due to an increasing urbanization and population growth in general [1]. growth occurs vertically, horizontally or in both directions [2]. vertical growing cities need more high-rise buildings to accommodate the rising number of people. because of sustainability aspects and numerous technical innovations at the turn of the millennium, timber multi-storey buildings came into focus [3]. currently, there are visionary timber high rise projects with building height up to 300 m [4]. the increasing building height leads to high forces in vertically load bearing members in the lower floors. this causes a mass consumption of material and space in timber-only high-rise buildings, due to a relative low compression strength and youngs modulus of timber compared to ultra-high-performance concrete. compared to the sustainable construction material timber, ultra-high-performance concrete may be regarded as not-sustainable, because of its high content on non-renewable material and its energy consumption for processing. aim is the sustainable use of ultra-highperformance concrete (uhpc) through its targeted use for vertically load bearing members. therefore, timber and ultra-high-performance concrete are combined in walls. the wall cross section is comprised of a slender uhpc core, encapsulated within timber, as shown in figure 1. applied normal forces concentrate on the concrete core. the surrounding timber prevents the slender concrete core from buckling and bears compression and tensile forces, caused by bending moments due to eccentricities or figure 1. wall with hybrid cross section comprised of timber and uhpc [5] external loads. the walls are economically produced in the industrialized environment of the cross laminated timber (clt) production. three steps are necessary. first, a precast concrete plant prefabricates lamellas. in the regular production process of a clt plant, the manufacturer inserts the prefabricated concrete lamellas into the raw clt panels at previously defined places. after pressing the panels, wall elements are formed out with joinery machines, automatic cutting machines or machining centres. to ensure statically monolithic acting walls and to enable this production process, a glued connection between timber and concrete using standard glues in 417 https://doi.org/10.14311/app.2022.33.0417 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en t. oberndorfer, f. hunger, o. fischer acta polytechnica ctu proceedings � figure 2. typical stress-strain relationship of different types of concrete according to [7] clt production is inevitable. 2. fundamentals 2.1. ultra-high-performance concrete ultra-high-performance concrete is characterized by its, for concrete, high compressive strength, up to 200 n/mm2, and a young’s modulus between 45.000 and 55.000 n/mm2, which is shown in figure 2. the high compressive strength comes along with a brittle failure when no steel fibres are added to the concrete. a certain degree of ductility can be achieved by adding steel fibres [6]. the high compressive strength of uhpc is achieved through its optimised packing density of the aggregate. therefore, almost no pores exist in uhpc. this leads to a very dense surface of uhpc compared to normal strength concrete and the possibility to easily glue uhpc. [8] 2.2. cross laminated timber cross laminated timber is a quasi-rigid composite, plate-like engineered timber product, which is commonly composed of an uneven number of layers, each made of boards placed side-by-side, which are arranged crosswise to each other at an angle of 90◦c [9]. the layers are normally glued together using polyurethane (pu) or melamine-urea-formaldehyde resin (muf) adhesives [10]. the thickness of clt elements is commonly between 51 mm and 300 mm, whereas the single layer thickness generally varies between 20 mm and 40 mm. common lengths of clt elements is around 18 m with a width up to 3,0 m [9]. figure 3 gives an overview of the clt production process and figure 4 displays a simplified technical drawing of a clt element. clt bears in-plane and out-of-plane loads through its crosswise layering. the load bearing behaviour is influenced by the cross wise layering and for stability and deformation calculations the shear deformation must be considered [11]. � figure 3. production process of cross laminated timber according to [9]. figure 4. simplified technical drawing of a clt element according to [9]. 2.3. timber concrete composite a well-known timber concrete composite (tcc) member is the tcc slab. in this composite element the concrete is positioned in the compressive zone, whereas the timber is placed in the tensile zone. the two elements are connected with special connectors to form one statically acting cross section. the effectiveness of the tcc element is mainly characterized by its connection. the usage of systems with high composite action can reduce beam depth and allow for longer span length [12]. adhesives are a very effective connector for tcc elements and lead to a quasi-rigid connection which increases the member stiffness and strength. in addition, the shear forces are distributed uniformly over the entire surface and local concentrations are avoided. [13] in the past, lots of research on glued connections has been performed, e.g. [14–17], most of these research projects have the use of epoxy resin as adhesive in common. 3. small scale experiments 3.1. general in the beginning of the project numerous small-scale experiments were performed to determine the possibility to glue timber to uhpc using standard glues for clt production. this was an iterative process, as the information obtained served as basis for 418 vol. 33/2022 glued timber concrete composite walls figure 5. experimental test setup for shear experiments [5]. the following experiments. these small-scale experiments were shear, delamination and bond experiments, which will be discussed in this section. 3.2. shear first, small-scale shear experiments on bonded timber concrete composite specimen were performed to find combinations of concrete, concrete surface treatments and adhesives with a promising behaviour. the three-layered specimen consisted of a concrete layer, placed in the middle, and two layers of spruce, which were free of any defects. per specimen two shear joints existed and shearing off took place parallel to the grain. three different types of concrete (c1, c2 and c3) were tested with and without steelfibres. c1 and c2 were commercially available compounds and c3 was a laboratory mixture. the used concrete surfaces were a) smooth, directly out of the formwork, b) ground and c) sandblasted. five adhesives were used: a one-component polyurethane adhesive, a melamine-urea-formaldehyde resin, a phenol resorcinol resin, a two-component polyurethane resin and an epoxy resin. the adhesive application was 400 g/m2 and the pressure at bonding was about 0, 75 ± 0, 1 n/mm2 for all specimen. the test setup is displayed in figure 5. a combination of concrete, concrete surface treatment and adhesive was characterized as promising if it achieved high shear strength and timber fracture in the experiment. considering only standard clt production adhesives, suitable results could be achieved independently from the used concrete type, with a ground concrete surface in combination with the onefigure 6. technical drawing of experimental test setup of inclined bond tests [5]. component polyurethane adhesive. specimen glued using the melamine-urea-formaldehyde resin failed in the glue line, which was characterized through a low shear strength and a maximum timber fracture of 20%. 3.3. delamination although this research project concentrated on the short-term load bearing behaviour of clt walls with glued in lamellas from uhpc, durability keystroke tests with glued specimen were performed. the test procedure included storing the specimen in boiling water for six hours, the storage in 20 ◦c cold water for one hour and afterwards drying in an oven for approximately 16 hours. this induces shearand tensile stresses on the glue line due to the swelling and shrinkage behaviour of the timber. if the glued connection is insufficient there will be a gap opening of the glue line. according to din en 16351 [18] the maximum tolerated failure of a single glue line is 40% of its length and 10% of glue line length may be opened for the sum of all glue line lengths. based on the shear experiments, delamination tests with promising combinations were performed. here, both steel-fibre reinforced concrete types c1 and c2, a smooth and ground surface, the one-component polyurethane adhesive, the twocomponent polyurethane resin, and the epoxy resin were used to form the specimens. the manufacturing procedure of the specimen was identical as for the shear tests. for each series two specimen were produced. 3.4. bond to verify the assumption of a rigid connection inclined shear tests with three different bond lengths, 419 t. oberndorfer, f. hunger, o. fischer acta polytechnica ctu proceedings figure 7. schematic (left) and real (right) experimental test setup for full scale tests [5]. 14 cm, 28 cm and 45 cm, were performed. a technical drawing of the experimental test setup is displayed in figure 6. the specimens were comprised of concrete type c1 lamellas with smooth and ground surface, glued to spruce glulam gl24c using the onecomponent polyurethane adhesive. bonding parameters were identical to the shear and delamination tests. all experiments showed a brittle failure when reaching the maximum force. with increasing bond length, the average shear force leading to failure increased. the average failure shear force for the 14 cm bond length with the ground concrete surface was 77,1 kn and increased up to 217,1 kn for the 45 cm long bond length. no difference could be seen in the behaviour of the smooth or ground concrete surface. the maximum measured relative slip between concrete and timber, recorded with an optical measuring system, in the middle of the bond length was approx. 0,25 mm for all experiments. in most cases the failure occurred in the timber part. only in two experiments, with the 14 cm long bond length, the failure occurred in the glue line. 4. full scale tests 4.1. methodology to analyse the load bearing behaviour of the composite walls, centrically and eccentrically loaded, experiments at scale 1:1 have been performed. figure 7 displays the test setup. for every eccentricity, 0 mm, 10 mm and 20 mm, three wall segments were tested. the tested wall segments were 0,5 m in depth, 2,84 m high and had a thickness of 15 cm. two roller tilting bearings were used to test a pendulum rod with a buckling length of 2,97 m. the five layered wall segments with a complete concrete core, every layer was 30 mm thick, were produced in a clt production facility. the adhesive application of a onecomponent polyurethane was production based 170 g/m2 and the bonding pressure was 0,08 n/mm2, because the elements were pressed in a vacuum bed. the used timber was spruce with a strength class of c24, the concrete had a mean compressive strength of fck, cyl = 132, 5n/mm2 and a mean young’s modulus of 46.774 n/mm2. these values were determined according to din en 12390-13 [19] and din en 123903 [20]. the tests were conducted displacement controlled with a speed of 0,12 mm/min. laser distance sensors controlled the horizontal displacement of the wall segments during the experiments. strain gauges recorded the strain on the outermost timber fibres and in the middle of the cross section on the concrete. the strain gauges were placed in the centre, at the bottom and the top of the wall. 4.2. results figure 8 displays the force-horizontal deflection in the vertical wall centre and the force-piston stroke diagram for the centrically loaded experiments. all three experiments showed similar behaviour and the maximum force varied between 1331 kn for v_d_30_00_3 and 1693 kn for v_d_30_00_2. with increasing load, the horizontal deflection increases slowly until the critical point is reached and the walls buckle. this is indicated by a rapid increasement of the horizontal deflection, accompanied with a decreasing load. after the specimen were removed out of the testing machine material failure could visibly not be detected. 420 vol. 33/2022 glued timber concrete composite walls figure 8. force-horizontal deflection and forcepiston stroke diagram for centrically loaded experiments [5]. figure 9. force-horizontal deflection and forcepiston stroke diagram for 10ămm eccentrically loaded experiments [5]. the force-horizontal deflection diagram of the 10 mm eccentric loaded experiments is shown in figure 9. the three specimens showed a similar behaviour upon reaching the maximum force, which varied between 908 kn and 1025 kn. the horizontal deflection at the maximum load is close to 30 mm. experiments v_d_30_10_1 and v_d_30_10_2 showed a ductile behaviour, whereas v_d_30_10_3 failed brittle. material damage could clearly be seen on the compressive side of the wall segments in the outer most timber lamella. the recorded strains in the wall centre showed a similar behaviour for all experiments. figure 10 shows a typical force-strain diagram, explicitly for experiment v_d_30_10_1. the concrete middle layer suffers only compressive strain, whereas the outermost timber fibres are either under compressiveor tensile strain. 4.3. analysis to calculate the maximum buckling force for the centrically loaded experiments, euler’s formula using an effective bending stiffness, accounting for the additional shear deformation, according to [11], was used. the eccentric experiments were analytically modelled figure 10. development of strains at the vertical wall centre for timber and concrete for experiment v_d_30_10_1. [5]. figure 11. development of strains at the vertical wall centre for timber and concrete calculated (gray) and experiment (black). [5]. with the differential equation for second order problems, also using an effective bending stiffness. due to the scattering of the elastic timber properties the mean stiffness values and the characteristic material strength of nominal timber strength class c24 according to din en 338 [21], and for concrete the determined young’s modulus and compressive strength were used for the calculations. the calculated buckling forces were lower than the ones achieved in the experiments. figures 11 and 12 compare the calculated strain distribution and horizontal deflection for different timber strength grades in the vertical wall centre with ones obtained in an experiment. the curves fit well. because a linear elastic material model was used to model the experiments, the declining branch in the experiment cannot be considered. 421 t. oberndorfer, f. hunger, o. fischer acta polytechnica ctu proceedings figure 12. force-horizontal deflection calculated (gray) and experiment (black). [5]. 5. conclusions the completed test program showed that it is possible to bond timber to uhpc using standard glues for clt production regarding short-term behaviour. the tests on wall segments verified, that the cross section can be considered as monolithic and the elements can be calculated using an effective bending stiffness. the maximum load capacity depends on the slenderness of the walls and the eccentricities. compared to homogenous clt walls, the hybrid walls tested with a 10 mm eccentricity, show a 25% higher load bearing capacity, calculated with characteristic values for material strength. acknowledgements the authors want to thank the forschungsinitiative zukunftbau for funding this project. this paper is only a short introduction in the project, the described experiments, and results in section three, four and five are detailly described in [5]. references [1] united nations, department of economic and social affairs, population division. world urbanization prospects 2018: highlights (st/esa/ser.a/421), 2019. https://population.un.org/wup/publications /files/wup2018-highlights.pdf. [2] a. janser, g. aerni, k. gimm. hochhaus wunsch und wirklichkeit. ostfildern hatje cantz, pp. 163, 2011:. [3] h. kaufmann h, s. krötsch, s. winter. atlas mehrgeschossiger holzbau, münchen, 2018. [4] council on tall buildings and urban habitat. the skyscraper center, 2019. http://www.skyscrapercent er.com/quick-lists#q=&page=1&type=building&stat us=vis&structural_material=timber&min_year=0&ma x_year=9999&region=0&country=0&city=0. [5] t. oberndorfer t, f. hunger, o. fischer. ultra high performing timber walls einsatz von schlanken lamellen aus ultrahochfestem beton in brettsperrholzelementen zur steigerung der tragfähigkei. www.zukunftbau.de. [6] e. fehling, m. schmidt, j. walrace j, et al. ultrahigh performance concrete uhpc (betonkalender 2013), berlin: wilhelm ernst & sohn, 2014. [7] t. leutbecher. rissbildung und zugtragverhalten von mit stabstahl und fasern bewehrtem ultrahochfestem beton (uhpc), kassel: kassel university press, 2008. https://kobra.uni-kassel.de /handle/123456789/2009031226631#. [8] b. sagmeister. maschinenteile aus zementgebundenem beton, berlin: beuth, 2017. [9] r. brandner, g. flatscher, a. ringhofer, et al. cross laminated timber (clt): overview and development. european journal of wood and wood products 74(3):331-51, 2016. https://doi.org/10.1007/s00107-015-0999-5. [10] p. mestek p, n. werther, s. winter. bauen mit brettsperrholz ed studiengemeinschaft holzleimbau e.v., wuppertal:studiengemeinschaft holzleimbau chapter 1, 2010. [11] s. winter, h. kreuzinger, p. mestek. holzbau der zukunft. teilprojekt 15. flächen aus brettstapeln, brettsperrholz und verbundkonstruktionen, stuttgart:fraunhofer irb, 2009. [12] d. yeoh, m. fragiacomo, m. de franceschi, et al. state of the art on timber-concrete composite structures: literature review. journal of structural engineering 137(10):1085-95, 2011. https: //doi.org/10.1061/(asce)st.1943-541x.0000353. [13] a. dias a, j. schänzlin, p. dietsch. design of timberconcrete composite structures: a state-of-the-art report by cost action fp1402/wg4, aachen: shaker, 2018. https://www.cost.eu/uploads/2018/11/design-of-t imber-concrete-composite-structures.pdf. [14] j. h. j. d. o. negrão, c. a. leitão de oliveira, f. m. maia de oliveira, et al. glued composite timber-concrete beams. i: interlayer connection specimen tests. journal of structural engineering 136(10):1236-45, 2010. https: //doi.org/10.1061/(asce)st.1943-541x.0000228. [15] e. ferrier, p. labossière, k. w. neale. mechanical behavior of an innovative hybrid beam made of glulam and ultrahigh-performance concrete reinforced with frp or steel. journal of composites for construction 14(2):217-23, 2010. https: //doi.org/10.1061/(asce)cc.1943-5614.0000063. [16] kanócz j and bajzecerová v 2015 timber concrete composite elements with various composite connections part 3: adhesive connection wood research 60 939-52. [17] m. schäfers, w. seim. investigation on bonding between timber and ultra-high performance concrete (uhpc). construction and building materials 25(7):3078-88, 2011. https: //doi.org/10.1016/j.conbuildmat.2010.12.060. [18] din en 16351:2015-12. holzbauwerke brettsperrholz anforderungen, 2015. 422 https://population.un.org/wup/publications/files/wup2018-highlights.pdf http://www.skyscrapercenter.com/quick-lists#q=&page=1&type=building&status=vis&structural_material=timber&min_year=0&max_year=9999&region=0&country=0&city=0 www.zukunftbau.de https://kobra.uni-kassel.de/handle/123456789/2009031226631# https://doi.org/10.1007/s00107-015-0999-5 https://doi.org/10.1061/(asce)st.1943-541x.0000353 https://www.cost.eu/uploads/2018/11/design-of-timber-concrete-composite-structures.pdf https://doi.org/10.1061/(asce)st.1943-541x.0000228 https://doi.org/10.1061/(asce)cc.1943-5614.0000063 https://doi.org/10.1016/j.conbuildmat.2010.12.060 vol. 33/2022 glued timber concrete composite walls [19] din en 12390-13:2014-06. prüfung von festbeton teil 13: bestimmung des elastizitätsmoduls unter druckbelastung (sekantenmodul), 2014. [20] din en 12390-3-2019:10. prüfung von festbeton teil 3: druckfestigkeit von probekörpern, 2019. [21] din en 338: 2016-07. bauholz für tragende zwecke festigkeitsklassen, 2016. 423 https://doi.org/10.14311/app.2022.33.0193 acta polytechnica ctu proceedings 33:193–199, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague energy flows along the production and use of secondary materials with a special focus on concrete karin gruhler∗, georg schiller leibniz institute of ecological urban and regional development, weberplatz 1, 01217 dresden, germany ∗ corresponding author: k.gruhler@ioer.de abstract. urban mining in the existing building stock can contribute to securing raw materials and conserving natural resources if the potential of recycling construction waste is consistently exploited. from an ecological point of view, it is on the one hand interesting what amounts of primary materials can potentially be substituted and on the other how much energy need to be invested for this. at present, the recycling of construction waste usually is considered from a material perspective. there is lack of an approach, extending material-oriented considerations by energetic aspects. the aim is to develop a uniform research approach by which energy expenditure during recycling of important construction products can be determined. besides concrete seven further construction products are investigated. recycling paths are described and analysed along waste management processing steps taking into account the quality of the demolition materials and the quality requirements of the possible new application variants in the construction sector. the result is a clear plea for more consistent recycling. the analyses of concrete indicate that "high-quality" recycling only results in energy advantages when "high-grade" demolition material is used. however, so-called "down-cycling" solutions allow resource conservation to be combined with energy savings, even with lower-quality demolition materials. the single-minded focus on "high-quality" recycling according to the general understanding should therefore be questioned. instead, preference should be given to solutions that take resource conservation into account in a more holistic way especially with regard to resource conservation and climate protection. keywords: circular economy, energy consumption, recycling, resource efficiency. 1. introduction the recycling of construction waste can make an important contribution to climate and nature protection, but above all to the conservation of resources. in germany, the german resource efficiency programme formulated the goals of decoupling resource consumption from economic growth, doubling raw material productivity by 2020, reducing the environmental impacts associated with the use of natural resources as far as possible and further developing and expanding the circular economy [1]. in order to enable almost circular economic activity in the construction sector, the corresponding demolition materials must be recyclable and be available with a certain consistency in specific quantities. resource conservation potentials using secondary materials expressed in tonnes of material are known for some selected construction products [2, 3]. however, an assessment of the resource conservation potentials only on the basis of these material figures is too one-sided, as energy aspects are not taken into account. there are individual studies that focus on specific construction products with regard to the energy used in recycling [4]. however, there is no comprehensive overview for all important construction products. there is a lack of a common approach by which important construction products can be presented synoptically and compared to each other in terms of their energy expenditures during recycling. the aim of the present study was to extend the material-oriented studies on resource conservation potentials to include energy considerations and to develop a uniform approach by which important construction products can be presented and compared in a synoptic way with regard to their energy expenditure during recycling [5]. the following questions should be clarified: which energy expenditures are associated with recycling? what is the energy expenditure of recycling compared to the energy expenditure of standard production without recycling? are there energetic differences between "high-quality" application variants (in building construction) and less "high-quality" application variants (in civil engineering and landscaping)? the investigations focused on eight construction products: concrete, bricks, sand-lime bricks, gypsum, flat glass, mineral (stone) wool, pvc profiles, and pvc floor covering. for all of these, we energetically investigated the paths from demolition material to a new application variant in building construction, civil engineering or landscaping. this we did taking into account the quality of the demolition material and the quality requirements of the new application 193 https://doi.org/10.14311/app.2022.33.0193 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en karin gruhler, georg schiller acta polytechnica ctu proceedings process chain demolition material secondary material application variant concrete pc 1 concrete without recycled aggregatescoarse adhesions 2/16 type 1 concerete foundation pc 2 recycled aggregates c20/25broken concrete mixture 2/22 type 2 pc 3 (with brick, limestone, plaster) recycled aggregates ballast base layer0/32 for road construction for roads table 1. model-like pc for concrete. variant in the construction sector. 2. methodology in order to carry out these investigations, we have developed an analytical approach which generally regulates the methodological procedure, following the idea of a continuous material flow analysis as described in [6]. 2.1. balance framework and terms & definitions starting point of the considerations are demolition materials. demolition materials are quantities that are available in a certain quality after the demolition of a building (e.g. broken concrete without coarse adhesion). these materials come exclusively from building construction. construction site waste is not included in the analysis. in order to be able to supply demolition materials as secondary materials and substitute materials for a new application variant, they must be prepared accordingly. a new application variant is the use of demolition materials for a new purpose and location (e.g. broken concrete as aggregates for foundation concrete). all necessary preparation and further processing steps depend on the quality of the demolition material itself as well as the quality requirements of the new application variant. in order to take these dependencies into account, there is a need to consider the entire process from the demolition material to the new application variant. this is applied in the form of characteristic model-like process chains (pc). due to the large variety of possible process chains (large variety of possible deconstruction qualities, types of processing, further processing methods, application variants) two to three demolition materials and two to three typical application variants were combined into model-like pc for each construction product. the pc are defined in consultation with experts from the relevant construction product associations in germany. table 1 shows the pc for the example concrete. on the way from the demolition material to the new application variant, a secondary material is first created and in a further step a substitute. the terms "secondary material" and "substitute" are described more in detail in the standard din en 15804:201407 [7] (table 2). with regard to din en 15804:2014-07, the selected balance framework starts at the end of the waste treatment step of the preceding product system "building". processes before the end of the waste property stage (such as collection and transport) are part of the waste treatment of the above product system "building". in contrast, processes that are required after the end of waste property stage to allow primary materials to be replaced in another product system are considered to be outside the above system "building" (see [7], para. 6.3.4.5 note 3). these regulations are complied with here. the analyses start with the demolition material at the recycling company and integrate all preparation steps for the production of the secondary material as well as its further processing up to the substitute (preceding demolition technology, collection processes, sorting processes and transports on the site are not considered). 2.2. 3-step analysis approach of process chains the energetic analyses of the model-like pc of the different construction products follow a uniform methodical approach. it is divided into three steps: 1. determination of the energy required to process a demolition material into a secondary material. 2. determination of the energy expenditure for the further processing of the secondary material into a substitute that can replace a primary material in a functionally equivalent way, using energy addons/deductions. 3. comparison of the energy expenditure for the substitute with that for the primary material to be replaced (figure 1). the steps of the methodical procedure are explained more in detail in the following paragraphs using concrete as example (pc 1: from pure crushed concrete to recycled aggregate for foundation concrete c20/25, table 1). 2.2.1. energy expenditure for processing secondary material here the energy required to process the demolition material into secondary material is calculated 194 vol. 33/2022 energy flows of secondary materials term description of important terms secondary material according to din en 15804:2014-07 [7], para 3.29, a secondary material is any form of material recovered from a previous use or from waste and replacing a primary material, recorded at the point at which the secondary material enters the system from another system, recovered from a previous use or from waste from a product system and used as input to another product system. substitute substitutes are "secondary materials having left the system for primary material production [and are declared] in module d if they have functional equivalence to the replaced primary material" ([7], para 6.3.4.5, note 3). table 2. description of the terms "secondary material" and "substitute" according to din en 15804:2014-07. figure 1. methodical procedure. in mj/kg. the following list shows which analysis and calculation steps are required to determine the average energy expenditure. • identification and definition of characteristic main processing steps to produce the secondary material (e.g. pre-screening, crushing, metal deposition etc.). • definition of technical components/machinery for the identified processing steps (e.g. sieve, jaw crasher, magnetic separator etc.). • determination of average energy expenditures for the defined components/machinery (e.g. sieve 0.0021 mj/kg, jaw crasher 0.0022 mj/kg, magnetic separator 0.0147 mj/kg etc.). • calculation of the energy required to produce the secondary material by adding the energy values of the identified process steps and defined machines (e.g. 0.036 mj/kg for all steps). • analysis of the material flows and calculation of the proportions for target product, co-products and waste. for pc 1 of concrete this is 48% target product (recycled aggregate 2/16 type 1), 43% coproducts (crushed concrete 16/22, 22/32, 32/56, 56/x and sand 0/1, 0/2) and 9% waste (pre-screen material 0/8 and impurities/foreign substances). for co-products there is the possibility of further use. they receive an energy allocation. there is no further use for waste and therefore no energy allocation. • calculation of the energy allocation for the coproducts based on the material flow proportions (e.g. 0.013 mj/kg for crushed concrete > 16 mm and sand ≤ 2 mm). • calculation of the energy required to produce the secondary material considering the energy allocations for co-products (e.g. 0.036 mj/kg − 0.013 mj/kg = 0.024 mj/kg) in sum step 1 ends with the energy expenditure for processing the secondary material. 195 karin gruhler, georg schiller acta polytechnica ctu proceedings figure 2. energy expenditures of the selected construction products a comparison. 2.2.2. energy expenditures for further processing up to a substitute here it is analysed which energy expenditure is connected with the further processing of the secondary material to the substitute. methodologically, all necessary measures and processes are analysed and the differences between the production process with recycled material and the standard production process without recycling are worked out. the differences can be associated with both additional and reduced energy expenditure. this means that the previous energy expenditure for the secondary material (0.024 mj/kg) receives an energy add-on or an energy deduction. additional and reduced energy expenditure result mainly from modified recipes and modified process steps1. modified recipes are, for example, different mixing ratios or additional admixtures of auxiliary materials. modified process steps can be additional, shortened or unnecessary process steps. in the case of concrete, it is above all necessary to increase the cement content in the recipe (+ 1%). modified process steps are not necessary (± 0%). this results in a total energy add-on of + 0.052 mj/kg. in total, this results in the energy expenditure for the substitute (0.076 mj/kg). 2.2.3. comparison of the energy expenditure for the substitute with that of the primary material finally, the production of construction products without recycled material is compared in terms of energy to that with recycled material. this is done 1modified transport efforts also lead to additional or reduced energy expenditure. however, they could not be calculated due to the poor availability of data. there is a need for research. at the level of the material to be replaced or the material to be substituted. in this way, the energy expenditure of the substitute is compared with that of the primary material. this comparison allows assessments to be made as to whether recycling makes sense from an energy perspective. the energy characteristics of the primary materials are taken from the source ökobaudat [8] and din en 15804 compliant secondary sources [9], [10], [11]. 3. results following the methodological procedure presented, energy expenditure for all pc of the selected construction products was calculated for the production of secondary materials and further processing up to the substitutes and compared with the energy expenditures of primary materials (to be replaced) (figure 2). in general, the results show that under the given framework there is almost nothing to prevent recycling. the energy expenditure for the secondary material and the substitute is usually lower than the energy expenditure for the primary material. exceptions are mineral wool and one pc of concrete (pc 2) and one of gypsum (pc 1). for mineral materials, recycling is not associated with excessive energy expenditures and makes sense above all with regard to the mass aspect, as recycling reduces the extraction of raw materials and protects the natural environment and landscape. for plastics the energy aspect is more significant. a specially look at concrete shows that the quality of the demolition material as well as the material requirements of the new application variant have a significant influence on the energy expenditures for recycling (table 3). as a rule, the prin196 vol. 33/2022 energy flows of secondary materials construction product concrete pc 1 pc 2 pc 3 demolition material concrete without broken concrete mixture coarse adhesions (with brick, limestone, plaster) ee of all processing steps 0.036 mj/kg 0.031 mj/kg 0.025 mj/kg discharge co-products 43% 34% 10% energy allocation (ea) co-products 0.012 mj/kg 0.009 mj/kg 0.002 mj/kg secondary material recycled recycled ra 0/32 aggregates (ra) aggregates (ra) for road 2/16 type 1 2/22 type 2 construction ee of all processing steps with ea 0.024 mj/kg 0.022 mj/kg 0.023 mj/kg substitute modified recipes + 0.052 mj/kg + 0.092 mj/kg ± 0 mj/kg modified process steps ± 0 mj/kg ± 0 mj/kg ± 0 mj/kg ee with add-ons/deductions 0.076 mj/kg 0.114 mj/kg 0.023 mj/kg for modification (77%) (118%) (14%) primary material aggregate of aggregate of aggregate of gravel (2/32; 80%) gravel (2/32; 80%) gravel round (4/x; and chippings and chippings 50%) and gravel (2/15; 20%) (2/15; 20%) broken (4/x; 50%) ee for processing primary 0.0972 mj/kg 0.0972 mj/kg 0.168 mj/kg material1 (100%) (100%) (100%) application variant concrete crushed stone foundation c20/25 sub-base (roads) 1 source: ökobaudat [8]. 2 using ecoinvent [9] results in a different figure: 0.057 mj/kg. this also influences the figure of the substitute due to the recipe. table 3. energy expenditure (ee) calculations of the different pc of concrete. ciple applies that "high-quality" demolition material (clean/without adhesions) is usually associated with a "high-quality" application variant (building construction) and a "low-quality" (unclean/ with adhesions) one with a "low-quality" application variant (civil engineering/landscaping). however, one and the same demolition material can also be used in different application variants (pc 2 and 3). for example, the demolition material crushed concrete mix can be used both as aggregates for the production of foundation concrete (pc 2) and as a ballast base layer in road construction (pc 3). the first variant has higher quality requirements and is also associated with higher energy expenditure. these result from a different secondary material: recycled aggregates 2/16 type 1 for concrete foundation (pc 2) and recycled aggregates 0/32 for road construction (pc 3). in addition, energy-relevant recipe adjustments (e.g. modified ratios for superplasticizer, water or cement) are required (pc 2) to produce a substitute equivalent to the primary material. a higher cement ratio content in particular has an impact on energy expenditure. compared to the two pc, the crushed stone sub-base for road is thus the energetically better application variant. likewise, two different demolition materials can have one and the same use (pc 1 and 2). in this way, both broken concrete without adhesions (pc 1) and crushed concrete mix (pc 2) can be used for the production of foundation concrete. here, however, the "low-quality" demolition material (crushed concrete mix) is the less efficient variant in terms of energy. it is true that the production of aggregates 2/22 type 2 requires less energy than that of aggregates 2/16 type 1. however, changes in the recipe finally lead to a higher energy expenditure. in comparison of the two process chains, the broken concrete without adhesions is therefore the more energy-efficient starting demolition material. the examples for concrete show that the use of recycling material in building construction can lead to different energy results with different types of demolition material (pure concrete or concrete mix). for example, the energy expenditure for the substitute in pc 1 is approx. 20% lower and in pc 2 approx. 20% higher than the energy expenditure for the primary material to be replaced (aggregates for concrete foundation). in contrast, energy savings of approx. 85% 197 karin gruhler, georg schiller acta polytechnica ctu proceedings are possible in road construction if recycled aggregates are used instead of primary aggregates (pc 3). if these results are taken into account, both "highquality" and "lower-quality" recycling should be given attention in the interests of resource and climate protection. 4. summary and conclusions in line with the objectives we developed a standard uniform balance approach to assess energy expenditure. using this approach we calculated two to three pc for each of the selected construction products und provided a first set of energetic figures for recycling. the analyses show that recycling is worthwhile. compared to the use of primary materials, it is generally not associated with an excessive use of energy. however, there are some exceptions. every building product has its own specific quality requirements and must be considered individually. for mineral materials, recycling makes sense above all with regard to the mass aspect, as recycling reduces the extraction of raw materials and protects the natural environment and landscape. for plastics the energy aspect is more significant. it also becomes clear that the pc of the construction products have to be considered from the demolition material until the new application variant is reached. this is because each new application variant is associated with certain quality requirements for the secondary material up to the substitute, which finally affect the energy expenditure. in summary, the analyses of concrete show that "high-quality" recycling (building construction application) only brings energetic advantages if "highquality" demolition material (pure concrete breakage) is used. "low-quality" recycling (road work application) on the other hand allow resource conservation to be combined with energy savings, even with lowerquality demolition materials. the single-minded focus on "high-quality recycling" according to the general understanding should therefore be questioned. instead, preference should be given to solutions that take resource conservation into account in a more holistic way especially with regard to resource conservation and climate protection. in order to be able to assess this comprehensively, it is always necessary to look at the entire process chain, starting with the construction waste and its quality, the intended recycling product and the intermediate treatment and processing steps. this should not stop at today’s common technologies. rather, it is important to keep an eye on innovative technical developments. for example, the use of geopolymers as cement substitutes can lead to a significant reduction in greenhouse gas emissions [12] and thus have an impact on the advantages of the recycling processes under consideration. acknowledgements the authors would like to thank the federal institute for research on building, urban affairs and spatial development (bbsr) for funding the research as well as the entire project team. in particular, we would like to thank tamara bimesmeier, who made most significant contributions to the development of the results of this publication. references [1] bmub. deutsches ressourceneffizienzprogramm ii: programm zur nachhaltigen nutzung und zum schutz der natürlichen ressourcen, p. 144, 2016. https://www.bmu.de/fileadmin/daten_bmu/pools/ broschueren/progress_ii_broschuere_bf.pdf. [2] g. schiller, c. deilmann, k. gruhler, et al. ermittlung von ressourcenschonungspotenzialen bei der verwertung von bauabföllen und erarbeitung von empfehlungen zu deren nutzung, dessau-roßlau: umweltbundesamt, uba-texte 56/2010, p. 195, 2010. https://www.umweltbundesamt.de/publikationen/ ermittlung-von-ressourcenschonungspotenzialen-bei. [3] c. deilmann, n. krauß, k. gruhler, et al. materialströme im hochbau. potenziale für eine kreislaufwirtschaft (zukunft bauen: forschung für die praxis), p. 88, 2017 https: //www.bbsr.bund.de/bbsr/de/veroeffentlichungen/ zukunftbauenfp/2017/band-06.html?nn=1143814. [4] m. buchert, j. sutter, h. alwas, et al. ökobilanzielle betrachtung des recyclings von gipskartonplatten. endbericht, dessau-roßlau: umweltbundesamt, uba-texte 33/2017, p. 110, 2017. https://www.umweltbundesamt.de/sites/default/ files/medien/1410/publikationen/2017-04-24_ texte_33-2017_gipsrecycling.pdf. [5] t. bimesmeier, k. gruhler, c. deilmann, et al. sekundörstoffe aus dem hochbau forschungsinitiative zukunft bau, band f 3184, p. 252, 2019. https://www.baufachinformation.de/ sekundaerstoffe-aus-dem-hochbau/fb/253180. [6] g. schiller, k. gruhler, r. ortlepp. continuous material flow analysis approach for bulk nonmetallic mineral building materials applied to the german building sector. journal of industrial ecology 21(3):673-88, 2017. https://doi.org/10.1111/jiec.12595. [7] din en 15804:2014-07: nachhaltigkeit von bauwerken umweltproduktdeklarationen grundregeln für die produktkategorie bauprodukt, 2014. https: //www.beuth.de/de/norm/din-en-15804/195229515. [8] ökobaudat. informationsportal nachhaltiges bauen, 2018. http://oekobaudat.de/datenbank/ browser-oekobaudat.html. [9] ecoinvent. the ecoinvent database (zurich), 2019. https://ecoinvent.org/the-ecoinvent-database. [10] ibu. published epds (institut bauen und umwelt), 2018 https://ibu-epd.com/en/published-epds/. [11] sphera. gabi solutions (ef database v2.0.). https://gabi.sphera.com/ce-eu-english/index/. 198 https://www.bmu.de/fileadmin/daten_bmu/pools/broschueren/progress_ii_broschuere_bf.pdf https://www.umweltbundesamt.de/publikationen/ermittlung-von-ressourcenschonungspotenzialen-bei https://www.bbsr.bund.de/bbsr/de/veroeffentlichungen/zukunftbauenfp/2017/band-06.html?nn=1143814 https://www.umweltbundesamt.de/sites/default/files/medien/1410/publikationen/2017-04-24_texte_33-2017_gipsrecycling.pdf https://www.baufachinformation.de/sekundaerstoffe-aus-dem-hochbau/fb/253180 https://doi.org/10.1111/jiec.12595 https://www.beuth.de/de/norm/din-en-15804/195229515 http://oekobaudat.de/datenbank/browser-oekobaudat.html https://ecoinvent.org/the-ecoinvent-database https://ibu-epd.com/en/published-epds/ https://gabi.sphera.com/ce-eu-english/index/ vol. 33/2022 energy flows of secondary materials [12] m. sandanayake, c. gunasekara, d. law, et al. greenhouse gas emissions of different fly ash based geopolymer concretes in building construction. journal of cleaner production 204:399-408, 2018. https://doi.org/10.1016/j.jclepro.2018.08.311. 199 https://doi.org/10.1016/j.jclepro.2018.08.311 https://doi.org/10.14311/app.2022.33.0119 acta polytechnica ctu proceedings 33:119–124, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague effect of superplasticizer on the strength of fly ash based geopolymer concrete rhem leoric cantos dela cruz∗, leah monica alaurin roriguez, emmanuel josh yambing tiongco, khim denize asilo yulas, jason maximino ongpeng de la salle university, gokongwei college of engineering, department of civil engineering, 8/f bro. andrew gonzalez fsc hall, 2401 taft avenue, 1004 manila, philippines ∗ corresponding author: rhem_delacruz@dlsu.edu.ph abstract. in this research, the effect of a polycarboxylate based superplasticizer on the strength of geopolymer concrete (gpc) was investigated. a fixed amount of superplasticizer (1.5 % of fly ash weight) was utilized along with alkali activators sodium hydroxide (naoh) and sodium silicate (na2sio3). the ultrasonic pulse velocity test determined the strength development of concrete. the quality of the gpc improved when a higher concentration of alkali activator was applied based on the concrete quality designation. in detecting the color development of gpc, samples were put through matlab and specimens became lighter as time passes due to dehydration (a process where water escapes from the sample). stress strain diagrams were generated which generally indicate that gpc specimens are ductile. the researchers were able to assess the workability of the mix designs using a rating from 1 to 5, with 1 being the least workable and 5 being the most workable. keywords: alkali activators, color detection, fly ash based geopolymer concrete, matlab, polycarboxylate based superplasticizer, ultrasonic pulse velocity test. 1. introduction throughout the years, the construction industry has developed and changed, paving the way to new improvements and innovation, especially to construction materials such as concrete. concrete consists of cement, fine aggregates, coarse aggregates, water, and additives to aid in the preparation of the mix. approximately 86.3 million metric tons of cement was manufactured in the united states and 4,100 million metric tons were produced worldwide in 2017 [1]. as the world continues to develop, more infrastructures are being made to meet the demands and necessities of people, which results in the increased usage of concrete. unfortunately, concrete usage greatly affects the environment since it is known to contribute around 5% of total world carbon emissions [2]. in the production of concrete, ordinary portland cement (opc) is commonly utilized as the binder. though opc may be a commonly used binder, it also has its downsides. a considerable amount of natural resources is depleted in order to produce ordinary portland cement [3]. the opc production emits large amounts of carbon dioxide. it contributes around 1.35 billion tons yearly or around 7% of the total greenhouse gas (ghg) emissions to the atmosphere of the earth [4]. cement production also requires a considerable amount of energy. in response to this, actions have been done to lessen the carbon emission. in the study regarding the reduction of the concrete industry’s impact on the environment, [5] suggested that less materials and energy should be used, and carbon dioxide emissions should be lessened. in order for the concrete to be more ecological, the opc used must be replaced with other binders. alternative binders such as rice-hull ash (rha) and fly ash (fa) are currently being used and tested to lessen carbon emission while maintaining the needed mechanical properties of a concrete mix as per astm. additionally, [6] presented in their review that it is feasible to reduce up to 80% to 90% of the current co2 emissions globally through the use of geopolymer technology. apart from reducing greenhouse gas emissions from the cement production, it also mitigates the issue of solid waste. geopolymer, an inorganic alumino silicate polymer made from secondary product materials which contain silicon and aluminum, can be utilized as an alternative to cement paste in the production of concrete [4]. geopolymer concrete undergoes the process of geopolymerization, similar to the hydration of cement. geopolymer concrete requires by-products with high aluminum and silicate content to react with an activator. different by-products may be utilized as a material for geopolymer binders, such as fly ash. fly ash, a by-product generated from coal fired power production, can be an alternative for cement. it is considered the fifth largest raw material resource in the world as 500 million tons of it, which is about 75% − 80% of the total ash, are produced annually [7]. fly ash is considered as an environmental pollutant when released into the atmosphere, however 119 https://doi.org/10.14311/app.2022.33.0119 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en r. l. c. dela cruz, l. m. a. roriguez, e. j. y. tiongco et al. acta polytechnica ctu proceedings composition (%) component point 1 point 2 point 3 average calcium (ca) 7.11 7.13 7.07 7.10 silicon (si) 22.37 22.37 22.24 22.33 aluminum (al) 9.34 9.5 9.58 9.48 table 1. xrf result of coal fly ash. there are numerous areas wherein fly ash can be applied, such as cement clinkers, road basement material, waste stabilization or solidification, concrete production, and geopolymer concrete [8]. as an additive to the cementitious material, fly ash decreases heat of hydration, reduces thermal cracking of concrete in early stages, improves mechanical and durability at late ages, and improves workability. a research conducted by [9] presented the different properties of a fly ash based geopolymer concrete and concluded that fly ash based geopolymer concrete has a compressive strength that is of high quality and can be used for construction purposes, and that its elasticity is similar to that of opc concrete, has excellent resistance against sulfate and acid, and undergoes minimal creep and drying shrinkage. despite the numerous research involving geopolymer concrete in the present time, studies on the utilization of admixtures in improving the strength of geopolymer concrete is still fairly limited. there is a need to further investigate the effect of the utilization of superplasticizer in gpc, as well as the impact of the curing methodology to be used in order to investigate it as the binder for concrete in construction. the objective of this study is to determine the optimum strength and workability of geopolymer concrete (gpc) with superplasticizer through the use of both mechanical and non-destructive testing. nondestructive testing in concrete has been advancing using nonlinear ultrasonic test [10, 11] and acoustic emission test [12, 13]. in this paper, ultrasonic pulse velocity test was used since it is practical and can estimate concrete strength [14]. 2. materials the study used an alumino-silicate coal fly ash obtained from a coal-fired power plant located in batangas, philippines. a hand-held x-ray fluorescence (xrf) spectrometer was used in order to acquire the elemental composition of the material, as per astm c188 shown in table 1. astm standards state that fly ash can be classified into three types: class f, c and n. commonly, class c and class f fly ash, which vary in terms of their calcium content, are used for the concrete binder. class c fly ash is high in calcium, while class f fly ash comprises less than 10% of calcium content [15]. observing table 1, it can be seen that the fly ash in this study contains low calcium and high silicon content. thus, the fly ash used in this study is a class f. a low calcium fly ash is favorable in making geopolymer concrete to obtain optimal binding properties [16]. sodium hydroxide also known as caustic soda flake; this alkali activator has a 98% minimum naoh purity. in this research, a molarity of 12m was used. in order to obtain such molarity, 480 grams of naoh pellets were gradually added to every 1 liter of distilled water. another alkali activator is the sodium silicate which is a colorless liquid. the said activator is known to be the most preferred activator for fly ash based geopolymer concrete [17]. the superplasticizer utilized in this research was a water-reducing admixture. its chemical base is a polycarboxylate ether with a white to yellowish color in appearance. the recommended consumption of the said admixture ranges from 0.8% 2.0% of the weight of the cementitious material. 15.51 20.62 12.99 19.25 0.00 5.00 10.00 15.00 20.00 25.00 1:1:1 1:1:2av er ag e c om pr es siv e s tre ng th , m pa design mixture without sp with sp figure 1. compressive strength of gpc without sp and gpc with sp. 3. methodology the study compared the gpc and gpc with superplasticizer with the aid of destructive and nondestructive tests. the destructive tests are astm c39 compressive strength test and strain gauge test. the nondestructive tests are ultra pulse velocity test (upv) and color detection test which were performed before and after the destructive testing. the following procedures were done upon mixing the raw materials: preparation of 12m naoh solution add 480 g of naoh pellets to every 1 liter of water; mix the dry materials such as fly ash, sand, and gravel; mix the wet materials such as sodium hydroxide (naoh) and sodium silicate (na2sio3); 120 vol. 33/2022 fly ash based geopolymer concrete materials (grams) ratio fly ash sand gravel na2sio3 naoh water-biner *sp (1.5% of fa) 1:1:1 1 1 1982.953 887.953 880.814 351.055 140.422 245.738 14.7 1:1:2 1 1 2737.215 665.964 1321.221 263.291 105.316 184.303 11 note: *sp is only for one set of specimens (those with superplasticizer) table 2. design mix of geopolymer concrete. design mix ratio rating description/remarks 1 : 1 w/o sp 3 difficulty in mixing was observed.the mixture can be mixed while exerting effort. 1 : 1 w/sp 5 the lightest and easiest to mix.the mixture easily spreads out. 1 : 2 w/o sp 2 notable lack of flow and relative dryness were observed.additional water was put so that it could be consolidated. 1 : 2 w/ sp 4 the mixture spreads out slowly with ease.the mix was dry before the superplasticizer was added. table 3. assessment of the workability of the geopolymer samples. figure 2. strength development of (a) geopolymer concrete without superplasticizer (b) geopolymer concrete with superplasticizer. figure 3. the concrete quality designation (cqd) derived from the upv. mixing of the dry and wet materials; add the water and add the superplasticizer. seen in table 2 is the design mix. 4. results and discussion 4.1. effect of superplasticizer on the workability workability, described as the ease and homogeneity of concrete during mixing, was one of the properties of concrete investigated in this study. the said fresh concrete property is reliant on several factors including water-binder ratio having its relationship as directly proportional. however, the strength of concrete is negatively affected as the water in the mixture increases. in gpc, alkali activators, superplasticizer, and water present in the mixture are the factors that affect the workability [18]. a high concentration of naoh and na2sio3 reduces the workable flow of geopolymer [19]. superplasticizer was used in this research to improve the workability of gpc. as expected, the mixture with superplasticizer resulted to be more workable than the mixture without super121 r. l. c. dela cruz, l. m. a. roriguez, e. j. y. tiongco et al. acta polytechnica ctu proceedings 0 5 10 15 20 25 30 0 1000 2000 3000 4000 5000 st re ss , m pa strain, μԑ 1:1:2 # 2 1:1:2 # 3 1:1:1 # 2 1:1:2 s # 1 figure 4. stress strain diagram of geopolymer concrete. figure 5. average color detection of each design mix. plasticizer. additionally, the researchers were able to assess the workability of various mix designs since all the specimens were hand mixed. shown in table 3 below is the assessment of the workability of the geopolymer in this study. the assessment makes use of a rating from 1 to 5, with 1 being the least workable and 5 being the most workable. 4.2. effect of superplasticizer on the compressive strength shown in figure 1 below is the comparison of the average compressive strength of geopolymer concrete without superplasticizer and that of the geopolymer containing superplasticizer. the figure 1 shows that the superplasticizer did not positively impact the compressive strength of the specimens. percentage differences of 17.68% and 6.87% were calculated for the design mixtures 1 : 1 : 1 and 1 : 1 : 2, respectively. the superplasticizer content of the mixtures was seen as the vital factor that affected its effect on the geopolymer concrete’s strength. increasing the content of polycarboxylate-based superplasticizer to 2% could positively impact the strength [20]. 4.3. strength development ultrasonic pulse velocity test the strength development of geopolymer was monitored in this research through ultrasonic pulse velocity (upv) test which was done every seven days until the 28th day. the velocity obtained from the upv machine was converted into its equivalent strength in megapascal (mpa) using raouf’s equation [21]. observing figure 2, it can be seen that there is a constant decrease in strength development for most of the samples. possible factors such as varying amounts of water and the inaccuracy of the manual compaction may have influenced the outcome of the results. based on the cqd by [22] shown in fig122 vol. 33/2022 fly ash based geopolymer concrete ure 3, the quality of the geopolymer concrete cylinders without superplasticizer and geopolymer concrete cylinders with superplasticizer were good and poor to questionable, respectively. 4.4. compressive strength test compressive strength test was done on the 28th day of the concrete cylinder samples. strain gauges were also attached to the sample upon the application of the load to generate stress-strain diagrams. table 4 shows the compressive strength of the geopolymer concrete samples. it can be observed that 1 : 1 : 2 without superplasticizer had the optimum compressive strength equal to 24.752 mpa. design mixture (fa:s:g) 28th day strength (mpa) 1 2 3 1 : 1 : 1 18.538 14.897 13.089 1 : 1 : 2 18.373 24.752 18.729 1 : 1 : 1 w/ sp 15.954 12.032 10.988 1 : 1 : 2 w/ sp 17.711 19.646 20.384 table 4. assessment of the workability of the geopolymer samples. figure 4 displays the stress strain diagram of the geopolymer concrete. it can be observed that 1 : 1 : 2 # 2 has the optimum compressive among all the specimens. the diagram also implies that the stressstrain curve of gpc is similar to a typical stressstrain curve of an ordinary portland cement concrete. additionally, the curve dictates the geopolymer is a ductile material. 4.5. image processing: color detection images were taken every 7th, 14th, 21st, and 28th day, which were then manually cropped to maintain a fixed perspective of the surface of the gpc. images were then uploaded to matlab so that the software can automatically convert them to grayscale, giving values between 0 (black) and 255 (white). the main factor in the change in color is due to gradual dehydration of geopolymer and oxidation of iron present in the fly ash [23]. each sample in figure 5 follows a pattern from a darker shade to a lighter shade, indicating gradual dehydration. 5. conclusion in determining the optimum strength of fly ash-based geopolymer concrete, the compressive strength test was used. the optimum mix design ratio would result in a compressive strength greater than 3000 psi (21 mpa). for gpc specimens, a peak strength of 24.752 mpa was obtained in the mix design ratio of 1 : 1 : 2. moreover, based on the previously shown stress-strain diagram, it can be stated that the geopolymer concrete followed a curved orientation until its failure point. such orientation denotes that the material tested is ductile. higher alkali activator concentration and proper method of mixing greatly affect the performance of geopolymer concrete. with the use of superplasticizer, the specimen had a glossier appearance when compared to a specimen without it. it was found that the superplasticizer did not increase the strength of fly ash-based geopolymer concrete. however, it improved the workability of the concrete specimens. further investigation is recommended on the application of different types of superplasticizers and other materials in improving workability of geopolymer concrete. references [1] t. wang. cement production globally and in the u.s. from 2010 to 2018 (in million metric tons). https://www.statista.com/statistics/219343/ceme 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[23] m. a. malik, m. sarkar, s. xu, et al. effect of pva/sio2 nps additive on the structural, durability, and fire resistance properties of geopolymers. applied sciences 9(9), 2019. https://doi.org/10.3390/app9091953. 124 https://doi.org/10.3390/app8030405 https://www.sid.ir/fileserver/je/103820160406.pdf https://doi.org/10.14359/51689238 https://doi.org/10.11113/aej.v7.15490 https://doi.org/10.1088/1757-899x/288/1/012019 https://doi.org/10.1109/natpc.2011.6136362 https://doi.org/10.1016/j.conbuildmat.2018.07.111 https://doi.org/10.1016/j.cemconcomp.2006.11.002 https://doi.org/10.1016/j.conbuildmat.2013.09.048 https://doi.org/10.1016/j.cemconcomp.2006.11.002 https://doi.org/10.1016/j.conbuildmat.2016.08.158 https://doi.org/10.3390/app9091953 https://doi.org/10.14311/app.2022.33.0160 acta polytechnica ctu proceedings 33:160–167, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague a new performance test to evaluate the sulfate resistance of concrete by tensile strength measurements volkert feldrappea, ∗, johannes haufeb, c, andreas ehrenberga, anya vollprachtb, thomas matscheib a building materials institute, bliersheimer str. 62, 47229 duisburg, germany b institute of building materials research, rwth aachen university, schinkelstr. 3, 52062 aachen, germany c zertplus überwachungsgesellschaft mbh, mühlenweg 11, 06749 bitterfeld-wolfen, germany ∗ corresponding author: v.feldrappe@fehs.de abstract. concrete structures without sufficient durability can be damaged by sulfates in groundwater and from surrounding rock layers. to evaluate the performance of a concrete mixture, precise and performance-oriented test methods are a must. therefore, a new a performance oriented concrete test procedure based on tensile strength measurements was developed considering experiences reported in international literature and recommendations of state-of-the-art reports. a vast parameter study with approx. 3850 tensile tests on astm briquets, 1900 flexural tensile tests on standard prisms and 2100 expansion tests on mortar flat prisms of different ages and with different storage conditions was statistically assessed. based on the results a performance-oriented test method could be defined which considers not only the chemical, but also the physical resistance of a concrete against sulfate attack. the method was verified by 23 concretes with different cements or cement fly ash combinations and additional field tests. it could clearly be demonstrated that the results represent the performance of a practical concrete in case of sulfate attack. furthermore, it leads much faster to an evaluation of the sulfate resistance compared to the most other practical oriented methods. keywords: concrete, performance test, sulfate attack, tensile strength. 1. introduction if concrete structures do not have a sufficient resistance, they can be damaged by sulfates dissolved in groundwater or incorporated in surrounding rock layers. such a sulfate attack is described as exposure class xa in the european concrete standard en 206. to ensure a sufficient concrete resistance, minimum requirements for concrete composition, such as a minimum binder content, a maximum accepted watercement ratio (w/c), accepted types of cement and additives as well as other protective measures, if necessary, are defined in en 206 together with their specific national concrete standards (e.g. din 1045-2 for germany). beside the descriptive measures the mechanisms that trigger damage because of sulfate attack have been extensively investigated in the past, e.g. in [1– 10]. a lot of test methods were developed worldwide. most of them can be used to assess the chemical resistance of cements and binders, respectively. only few, however, allow to test concrete. depending on the method, the test specimens are completely, partially or cyclically immersed in sulfate solutions. the concentration of the sulfate solution varies over a wide range, as does its temperature. in germany, only test methods are applied, which evaluate the chemical sulfate resistance of a binder. these are the socalled sva (sachverständigenausschuss) method of the dibt (deutsches institut für bautechnik), mainly used at present and the older methods by kochsteinegger [11] and wittekindt [12]. despite intensive research and optimization, all of them show considerable scattering of the results and test artifacts. these can be attributed, to extremely high sulfate concentrations of the test solution that are not related to practical applications. furthermore, the physical resistance the structural density of the concrete was deliberately neglected in these procedures. for these reasons, none of the methods has been included in german or european standardization so far [13, 14]. however, for sustainability reasons, it is a must to be able to evaluate the performance of a concrete in a precise and performance-oriented way, because new types of cements, additives and concretes with lower ecological footprints must continue to be developed in future. also, for technical or economic reasons, it is often advisable to deviate from the normative concrete specifications. moreover, there are current activities to shift the classical descriptive concept of concrete standardization to a performancebased concept. all these activities require a reliable test method that allows an unerring evaluation of the sulfate resistance of a concrete mix. 160 https://doi.org/10.14311/app.2022.33.0160 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 performance test for sulfate attack on concrete variable unit variation binder type − cem i; cem i-sr3; cem i +fly ash;cem ii/b-s; cem iii/a testing temperature ◦c 5, 12, 20 concentration of sulfate solution mg/l 3000, 6000 type of sulfate solution − na+, mg2+ as cation cement content kg/m3 320, 360, 400 content of cement + fly ash kg/m3 270 + 90, 285 + 94, 300 + 100 equivalent water-cement-ratio w/ceq 1 − 0.45, 0.50 1 w/ceq = w/(c + k · f ) with c = cement content, k = 0.4 and f = fly ash content table 1. influencing variables on the sulfate resistance and variation parameters. 2. objectives and realization the main objective of the research project was to develop a practical test method for the precise and selective determination of the sulfate resistance of concrete within an appropriate test period [15, 16]. the following three essential questions were defined to achieve the objective: 1. what are the test constraints that can be used to accelerate the damage mechanism without causing test artifacts? are the test parameters verifiable, and which damage can be recorded reproducibly and accurately? 2. is it possible to validate the results obtained with the new testing procedure with practical construction experience and field tests? 3. is there a limit value to differentiate between concretes with high and insufficient sulfate resistance? the study was based on experience gained from the currently applied methods as well as the findings of the state-of-the-art reports of dafstb (deutscher auschuss für stahlbeton) [13, 17, 18] and cen/tc 51 [14]. furthermore, both chemical resistance of the binder and physical resistance of the concrete structure were considered, since both partial resistances are important for the durability of concrete under practical conditions. therefore, parameters influencing the sulfate resistance of concrete were varied systematically at the beginning of the project. in addition, their effect on various parameters characterizing the microstructure was determined. the systematic statistical evaluation of the test results made it possible to define a test procedure based on the influencing parameters considered. this was followed by verification with another setup of approx. 25 concretes made with both binders known to have high and low sulfate resistance. at the same time, several of these concretes were stored under practical conditions at two different sites for at least one year [19]. the evaluation of all results made it possible to propose an acceptance criterion for the test procedure with which the sulfate resistance of a tested concrete can be evaluated reliably. 3. experimental setup 3.1. statistical design of experiments (doe) statistical methods of design of experiments were used intensively in order to consider a wide as possible test matrix. the influencing parameters considered in the statistical experimental design are summarized in table 1 [16]. in addition, their range of variation is also listed in the table. the statistical software minitab© was used for designing and analyzing the experiments. due to the large number of influencing parameters and the different verification levels, the experimental strategy was to create a full-factorial experimental design first and to select an optimal experimental design from it afterwards. for this purpose, individual combinations were chosen by doe software using methods of sequential optimization and taking into account terms up to the second order. the general full-factorial experimental design contained a total of 180 experiments covering all binders and a constant w/ceq at 0.45. the optimal design selected from it could be reduced to 100 experiments. based on the statistical analysis of the results, the original test design was adapted so that individual influencing parameters that did not show any significance were omitted and additional parameters that were deliberately not considered in the first step (e.g. w/ceq ) were included. a total of 120 tests series were considered in the parameter study for the development of the test method [16]. 3.2. concrete production and storage the binders listed in table 1 have been investigated in detail. the four commercially available cements were in line with en 197-1, the hard coal fly ash (fa) complied with en 450-1. their strength development measured according to en 196-1 is shown in table 2. concrete mixtures with those binders and quarzitic aggregate having a maximum grain size of 8 mm were produced. the w/ceq value followed the specifications of the optimal test design. the production of the so-called fine concretes was in accordance with en 12390-2. prisms with dimensions 40 × 40 × 160 161 v. feldrappe, j. haufe, a. ehrenberg et al. acta polytechnica ctu proceedings compressive strength after 2 d 7 d 18 d 91 d n/mm2 n/mm2 n/mm2 n/mm2 (2) cem i 42.5 n 28.5 46.6 61.9 71.0 (5) cem i 42.5 n-sr3 24.5 42.5 50.3 60.8 (14) cem ii/b-s 42.5 n 22.2 41.1 61.8 73.3 (15) cem iii/a 42.5 n 18.7 40.8 61.9 73.9 (21) fa 11 24.4 40.1 51.9 65.4 1 combination of 25 wt.-% fly ash and 75 wt.-% cem i 42.5 r. table 2. strength development of cements and fly ash. figure 1. concrete test specimens acc. to astm c307-03. mm3 for testing the flexural tensile strength and the dynamic modulus of elasticity and briquet specimens according to astm c307-03 (cf. figure 1) for testing the tensile strength were produced as test specimens. all specimens were demolded after one day and then stored in saturated ca(oh)2 solution at 20 ◦c for 27 days. 3.3. execution of tests and test parameters the test started at a concrete age of 28 days. the specimens were stored in sulfate solution at 5, 12 or 20 ◦c for 181 or 273 days according to the conditions defined in the doe. the flexural and tensile strengths as well as the dynamic modulus of elasticity were determined after 119, 181 and 273 days, where applicable. all individual values and no mean values were always used to relate them to respective reference values. the relative values obtained this way can be compared directly with each other. reference values were based on corresponding parameters obtained either on samples of same age but stored in saturated ca(oh)2 solution or determined before the start of sulfate storage. furthermore, a maturity function was used for the calculation of relative bending and tensile strengths. the basis was the function described in the fib model-code [20]. it was adapted to account for the influence of supplementary cementitious materials on the strength development of concrete, as proposed by vollpracht et al. [21]. since the adapted method was originally developed for the prediction of compressive strength, its suitability was tested in advance for bending and tensile strengths [15, 16]. by using single values and not mean values, the number of results for each individual experiment of the statistical design was increased. three single results each were used for the flexural tensile strength and the dynamic modulus of elasticity and even six single values for the tensile strength. this significantly increased the statistical certainty in the evaluation of the experimental designs. 4. results and discussion 4.1. suitable test criterions the different relative test parameters were analyzed in terms of their significance for the test procedure. damage to concrete due to sulfate attack was best characterized by the tensile strength of astm briquets. it was also shown that the common procedure, which uses the strength of specimens of the same age stored in a saturated ca(oh)2 solution as a reference value, gives comparatively wide scattering results. in contrast, tensile strengths based on the adapted maturity formula of the fib model-code show 162 vol. 33/2022 performance test for sulfate attack on concrete figure 2. contour diagram of relative tensile strength ft/ftm of concretes with (2) cem i 42.5 after 182 days storage in na2so4 solution, reference: fib maturation function. figure 3. mean decrease of the rel. tensile strength of the fine concretes with w/ceq = 0.50 with a change of the sulfate concentration of the na2so4 solution from 3000 mg/l to 6000 mg/l. significantly lower test scatter. the necessary test effort is also significantly reduced. therefore, this relative tensile strength ft/ftm is the appropriate test parameter for the new test procedure. 4.2. definition of the test procedure the evaluation of the parameter study provides statistically secured information on the significant influencing parameters and their contribution to the expected relative tensile strength ft/ftm. as an example, the evaluation is visualized in a contour diagram (figure 2) for fine concretes with portland cement (2) cem i 42.5 n after 182 days storage in na2so4 solution. it shows the effect of testing temperature and concentration of sulfate solution on the expected relative tensile strength. as a result, the lowest relative tensile strength can be expected if the concrete is tested at 5 ◦c and 6000 mg/l so 2−4 concentration. the figure also contains the results of six additional tests, carried out to verify the results of the statistical evaluation. each single significant influencing parameter was analyzed in terms of its effect on accelerating the testing and its potential tendency towards test artifacts. for example, figure 3 illustrates the accelerating effect when the sulfate concentration was increased. an increase of sulfate concentration from 3000 mg/l to 6000 mg/l led to a decrease of relative tensile strengths of comparable fine concretes between 0.10 and 0.15 (10 to 15 %) at a test age of 273 days, if damage occurred. at the same time, no excessive gypsum formation was observed in the pore 163 v. feldrappe, j. haufe, a. ehrenberg et al. acta polytechnica ctu proceedings concrete composition • fine concrete with max. aggregate size of 8 mm • w/ceq ratio 10 % higher than planned for concrete formulation • binder content as planned for concrete formulation test specimens • briquets acc. to astm c307-03 made from one concrete batch storage • 28 d in saturated ca(oh)2 at 20 ◦c test conditions • test solution: na2so4 • so42− concentration: 6000 mg/l • storage temperature: 5 ◦c • test duration: 273 days test parameters • relative tensile strength ft/ftm • ft: measured tensile strength at testing • ftm: tensile strength at testing, calculated by maturity function in accordance to fib model code • visual assessment (cracks, spalling, etc.) table 3. definition of the performance-oriented, test method for evaluating the sulfate resistance of concrete. figure 4. storage site (1) in a german gypsum mine. space of the specimens. consequently, the increase of sulfate concentration to 6000 mg/l accelerates the test and does not cause artefacts, especially for less sulfate-resistant concretes. as a result of this evaluation the test procedure could be described. its main features are summarized in table 3. 4.3. storage under practical conditions a large number of fine concretes as well as normal concretes the latter fulfilled the minimum requirements of din 1045-2 for the composition of exposure class xa2 were stored under practical conditions at two sites. figure 4 shows the exposure site in a gypsum mine. the laboratory results obtained with the new method will be verified by long-term tests with this real sulfate attack. many samples were stored over a period of more than one year and inspected in regular intervals. no concrete deterioration was detected during this period as it is shown in figure 5. it illustrates the relative dynamic modulus of elasticity of fine concretes stored in the gypsum mine. as expected, the storage time was too short to induce damage even on concrete that is known to possess insufficient sulfate resistance. the trials under practical conditions will be continued for some years. figure 5. rel. tensile strength of concretes after more than one year of storage under practical conditions. 4.4. proposal of an acceptance criterion after the definition of a new test procedure another task of the research project was to develop a proposal for an acceptance criterion for a reliable evaluation of the concrete sulfate resistance. for this purpose, 23 additional concretes with different cements and cement-fly ash combinations of different manufacturers were tested with the new test method. the relative tensile strength was determined after 119, 182 and 273 days. figure 6 illustrates the relative tensile strengths of the 23 concretes after 182 and 273 days of sulfate storage. after 119 days of storage no reliable statement can be made on the sulfate resistance. the first differences between concretes with different binders did occur after 182 days of storage. however, a definitive differentiation between concretes with known high or low sulfate resistance was not yet possible. some con164 vol. 33/2022 performance test for sulfate attack on concrete figure 6. relative tensile strength ft/ftm of concrete with 23 different binders after 182 and 273 days. cretes produced with portland cement without sr property showed residual tensile strengths comparable to those of some slag cement concretes, for which a high sulfate resistance can be expected in the light of experience. after 273 days (9 months) of sulfate storage, it was possible to make a clear distinction regarding the sulfate resistance of concrete. concretes with blast furnace cements cem iii/a or cem iii/b and portland cement/fly ash combinations obviously showed a high sulfate resistance with relative tensile strengths of 0.97 to 1.02. in contrast, concretes with portland cement including those with sr property and portland composite cements showed low residual tensile strengths and significant damage. the damage of concretes with cem i-sr cements was confirmed by further tests. c3a was determined by xray diffraction for all these cements. presumably, the c3a content is high enough to trigger a damaging ettringite reaction with the sulfate ions in the cement stone structure. considering the discussed results, it can be stated that a concrete has a sufficient sulfate resistance if its relative tensile strength is not lower than 0.70 (70 %) after 273 days of sulfate storage. furthermore, two stop criteria can also be defined for the test after 182 days. firstly, the test can already be stopped at this time if the relative tensile strength ft/ftm is lower than 0.70 (70 %), since the acceptance criterion will definitely not be reached even after 273 days. such a concrete will have a low sulphate resistance. secondly, the test can also be stopped at this point if the relative tensile strength ft/ftm is higher than 0.85 (85 %), because the acceptance criterion defined for a test age of 273 days is then also fulfilled with certainty. such a concrete will have a high sulfate resistance. 5. conclusions the focus of the research project was the development of a concrete test procedure based on tensile strength tests, which allows a clear differentiation between concretes with and without high sulfate resistance. in the development of the test method, the recommendations of the state-of-the-art report [13] 165 v. feldrappe, j. haufe, a. ehrenberg et al. acta polytechnica ctu proceedings were taken into account and the tensile strength was determined as the best test parameter [16]. based on the statistical evaluation of approx. 3850 tensile tests on astm briquets, 1900 flexural tensile tests on standard prisms and 2100 elongation tests on mortar flat prisms of different ages and after different pre-storage conditions, a new performance-oriented test method could be defined which was verified by 23 concretes with different cements or cement fly ash combinations. the concluding assessment of the research project is that the newly developed performance-oriented test method • can represent the performance of a practical concrete in case of sulfate attack, • considers not only the chemical, but also the physical resistance of a concrete against sulfate attack, • leads much faster to an evaluation of the sulfate resistance compared to common methods (current regulation sva test: testing at 3000 mg so42−/l and 5 ◦c for 2 years), • represents the damage mechanism more realistically than most conventional test methods and therefore leads to the avoidance of test artifacts, and • could also be carried out as a "binder test" if a fixed concrete formulation is used (e.g. the limit formulation of din 1045-2 for exposure class xa2). acknowledgements the igf project no. 19251 n of the research association vdeh-gesellschaft zur förderung der eisenforschung mbh was funded via the aif as part of the program for the funding of joint industrial research (igf) by the federal ministry of economics and energy based on a resolution of the german bundestag. the authors would like to express their thanks for the support. references [1] f. bellmann, b. möser, j. stark. influence of sulfate solution concentration on the formation of gypsum in sulfate resistance test specimen. cement and concrete research 36(2):358-63, 2006. https: //doi.org/10.1016/j.cemconres.2005.04.006. 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[21] a. vollpracht, m. soutsos, f. kanavaris. strength development of ggbs and fly ash concretes and applicability of fib model code’s maturity function a critical review. construction and building materials 162:830-46, 2018. https: //doi.org/10.1016/j.conbuildmat.2017.12.054. 167 https://doi.org/10.3390/cryst11091018 https://doi.org/10.1016/j.conbuildmat.2017.12.054 https://doi.org/10.14311/app.2022.33.0350 acta polytechnica ctu proceedings 33:350–356, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague comparing the environmental performances of new and renovated school buildings oscar mancinelli∗, alessandro pasquale fantilli, bernardino chiaia politecnico di torino, department of structural, building and geotechnical engineering (diseg), corso duca degli abruzzi, 24, 10129 torino, italy ∗ corresponding author: oscar.mancinelli@polito.it abstract. evaluating which is the best choice between renovating an existing construction or building a new structure in countries like italy, where a huge post-war un-listed building heritage does not satisfy the current standards and the economic resources are limited, is not trivial. several parameters come into play, such as such the extent of the construction work, the environmental cost of disposing old materials, the carbon footprint and volume of new materials. this paper is devoted to the analysis of two projects. the first consists of a renovation of a multi-storey existing school built in 1960s having total area of about 9900 m2. the second is a new construction of a three-story school having a total area of about 14000 m2 and made with timber. the results show that the existing school building, although having a lower embodied carbon related to materials, has a higher overall carbon footprint due to the co2 emissions related to operational energy. keywords: environmental sustainability, existing building, reinforced concrete, retrofitting, school buildings, timber. 1. introduction 1.1. state of art according to the latest report on the italian social situation [1], 30% of the italian school buildings were built before 1960, and 44% between 1960 and 1970. this implies that three-quarters of the schools have not been designed according to anti-seismic and energy-saving rules. besides, many of them show significant signs of degradation due to an insufficient maintenance. hence, actions aimed at improving and renovating the italian school building heritage are of fundamental importance and should be rapidly adopted. on the other hand, the environmental sustainability of the construction sector is a problem as well, as it needs a huge volume of raw materials and releases large quantities of greenhouse gases in the atmosphere. specifically, it accounts for 36% of global final energy use and 39% of energy-related carbon dioxide emissions, when upstream power generation is included [2]. also, construction and demolition waste (c&dw) is one of the most important waste streams, accounting for approximately 25% 30% of all the waste generated in the eu [3]. to solve all these problems related to the construction and management of school buildings, some strategies are necessary. the demolition of old and energy-intensive buildings and the construction of new low-energy (or even passive) building, made with materials with low carbon footprint, is definitely an efficient solution. in compliance with this strategy, in recent years, concrete has been replaced by more eco-friendly structural materials, such as wood [4], although studies proved that wooden structures show lower thermal performance than concrete, especially in long-term analyses [5], [6]. nevertheless, the demolition of old constructions, combined with the construction of new buildings, leads to the production of waste and to the consumption of natural resources to produce new materials. on the contrary, rehabilitation, refurbishment and renovation of old buildings make stream waste and the use of new materials limited. regrettably, in many cases, it is not always economically convenient to attain the same thermal and structural performances of a new building by adapting old building to the current code requirements. to provide guidance tools for policymakers and stakeholders, a large number of studies have been carried out, in which environmental performances of new buildings made of different materials are compared. as a result, in terms of global warming potential (gwp), timberframed residential buildings tend to have a lower environmental impact than those made with concrete and masonry [7]. regarding school buildings, studies on structural materials and on envelope systems have shown that concrete and masonry buildings have better thermal performances, because of the heavyweight materials [8]. although manufacture, construction and demolition of masonry and concrete buildings require larger energy and show higher global warming potential, these buildings exhibit lower annual energy consumption and environmental impact in service, which sometimes makes them more sustainable than those made with timber and steel [9]. based on these studies on refurbishment and newly constructed 350 https://doi.org/10.14311/app.2022.33.0350 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 environmental performances of new and renovated school buildings case study id gross surface(m2) estimated students type of school cost of construction (€) school building #1 sb #1 13990 1479 science school 23,500,000 school building #2 sb #2 9930 1002 technical school 10,600,000 table 1. main properties of the two schools. buildings, it is still not possible to conclusively determine which of the two alternatives has the best environmental performance over the entire lifespan [10]. besides, studies using the lca methodology mainly focus on energy refurbishment when the environmental impacts before and after intervention are compared. conversely, there are few lcas on the environmental impact of system reparation [11]. also, these approaches have to be reviewed, because the system boundaries of lca are not systematically explicit [12]. the lack of univocal procedures, as well as of the standardised methods of visualising lca results [13], leads researchers to apply methods and assumption that are appropriate for the single case study, making results difficult to compare. therefore, further investigations are needed both to finetune the assessment criteria and to broaden the assessment scenarios. 1.2. research significance is it more convenient to refurbish an existing building or to build a new school? to answer to this question, a benchmarking analysis on two case studies is proposed herein: • school building #1: a newly timber school building. • school building #2: an old school (built in 1960) in which energy efficiency and structural rehabilitation works have been carried out. 1.2.1. the two buildings the school building #1 is a 4-storey precast timber construction (one of which is the basement) with a total living area of about 14000 m2. this type of building has been chosen for the lower carbon footprint of the wood, as well as for the speed of construction given by the prefabrication. the building components (structures, envelope, internal partitions, etc.) fulfil the structural [14] and energy [15, 16] performances required by the current code rules for new buildings. besides, this school has been designed by implementing solutions aimed at optimising the use of climatic conditions and solar radiation, such as avoiding windows on the south side of the buildings and providing shading systems and large overhangs. accordingly, the building consists of a i-shaped main part, where classrooms, offices, laboratories and parking spaces are located. in a separate rectangular block, a sport hall is present. the school building #2 is a 6-storey building (including the basement) with a reinforced concrete frame and a total gross area of about 9900 m2. it consists of a multi-storey building, where mainly offices and classrooms are located, a two-storey building containing two sports halls and some classrooms, and a third single-storey block dedicated to teaching laboratories. refurbishment works, completed in 2013, aimed at complying the structural safety requirements provided by code rules for existing buildings [14]. specifically, the structure has been strengthened by applying steel cage systems for the columns (with l profiles and transverse plates), introducing steel braces, and strengthening the foundation by means of steel pipe piles. to meet the energy performances [15, 16], an extensive renovation of the building envelope was performed, by providing an insulation layer, installing ventilated facades and by substituting old windows and doors. table 1 summarises the main data of the two school buildings. due to confidentiality reasons, the name of the schools and the place where they are located are undisclosed. however, as they are in the same city, some boundary conditions are equal. these include the climatic conditions, the distance between the material production sites and the construction site, as well as the distance to landfill facilities. the life cycle assessment (lca) methodology is carried out to assess the environmental performance of both buildings. even if the environmental impact of the existing school before the renovation and the embodied carbon of the demolished materials are excluded. 2. materials and methods 2.1. methods the life cycle analysis carried out in this study is based on the jrc technical report [3]. the analysis is performed at production level, according to en 15804 [17], and at building level, in accordance with en 15978 [18]. in these standards, a modular approach for the definition of system boundaries is adopted, which enables to allocate the greenhouse gas emissions (ghgs) over the entire life of a building, i.e. from cradle-to-grave. it includes the building materials production (modules a1 to a3), the construction stage (modules a4 and a5), the use phase (modules b1 to b7) and the end-of-life phase (modules c1 to c4). lastly, module d considers the possible benefits and loads beyond the system boundary, namely those provided by the recycling, recovery and reuse of materials. the lifespan of both the buildings investigated herein is assumed to be 50 years, dur351 o. mancinelli, a. f. fantlli, b. chiaia acta polytechnica ctu proceedings figure 1. models created in designbuilder concerning: (a) sb #1; (b) sb #2. ing which the emissions related to maintenance and replacement of components are neglected. although assuming the same lifespan for both the refurbished and the construction of new building could not take into account the longer life expectancy of new school, it can be considered reasonable within the italian context. in fact, 50 years is the average life expectancy that italian building code assumes for ordinary structures [19], and it is also the average time that elapses between major refurbishment works [20]. 2.1.1. life cycle assessment both the schools have been modelled with revit software by implementing the stratigraphy and the geometric dimensions of the projects. in this way, the volumes of materials are computed and then used as input data to estimate the embodied carbon. the unitary impact of each material, expressed in accordance with the climate change indicator, is an input data. the global warming potential (gwp) indicator was selected for the assessment of the environmental impact of the two buildings, because it is consistent with the scope of the study. moreover, it is the most used indicator in the world, as it provides the result in terms of co2 equivalent, which is the main cause of global warming [21]. basically, there are two types of data sources from which the carbon footprint of materials can be obtained: generic databases (i.e., secondary data) and primary data, so-called environmental performance declarations (epds), provided by producers. results coming from the two categories of data can differ even of 25% for the gwp indicator [22]. this is due to the fact that secondary data are frequently unreliable, because they are based on average local data. conversely, epd is nowadays associated to all the new building material. thus, using product-specific primary data is recommended. in the current analyses, when no specific design indications were available, the same epd is associated with the common building components and materials (openings, insulation layers, etc.), in order to make the results of the two buildings comparable. 2.1.2. thermal analysis the evaluation of the co2 eq. emitted by the two buildings in 50-years of use (module b) has been carried out through the thermal analysis of the building, developed within designbuilder [23]. these models are shown in figure 1. as a result, the energy for heating and cooling the schools, and that used for domestic hot water (dhw) over the years is computed and converted in terms of co2 eq. by means of suitable conversion factors [24]. specifically, an electric heat pump system (eer = 1,75) is used to cool both the schools, whereas the needs for heating and dhw are produced by a heating system (cop = 0.85) powered by natural gas. system losses are not assessed. the emissions related to internal furniture and lighting system are neglected as well, because they are not comparable to the energy intensive laboratories present in sb #2. the overall impacts of the two buildings are calculated by summing the contributions from the shortterm (i.e. construction phase, in section 2.1.1) and the long-term (i.e., use phase, in section 2.1.2), assessments. to compare the results, two functional units are considered herein: • co2 eq. emissions per unit of gross building area (kg co2 eq/m2); • co2 eq. emissions per student (kg co2 eq/pers.) 2.2. materials for both the schools, information on the materials is provided, even if suppliers and location are unknown. thus, epd was selected for each material, assuming the same producer when it is present in the two buildings. on the other hand, the embodied carbon of the materials demolished in the school building #2 has not been taken into account, because the type, the amount, and the percentage of possible recycling, renewal or reuse are not included within the project. building materials are aggregated into three categories, namely skin, space plan, and structure, following the layered division (the so called 352 vol. 33/2022 environmental performances of new and renovated school buildings figure 2. embodied carbon of building materials grouped into macro-layers: (a) total amount; (b) amount per unit of gross building area (u.a.); (c) amount per student. figure 3. comparison of operational energy of buildings grouped by districts: (a) total amount; (b) amount per unit of gross building area; (c) amount per person. figure 4. comparison of overall co2 eq. emissions marked by sources (material and energy needs): (a) total amount; (b) amount per unit of gross building area; (c) amount per person. 353 o. mancinelli, a. f. fantlli, b. chiaia acta polytechnica ctu proceedings shearing layers) originally suggested by brand [25], and updated by other authors [26, 27]. through this approach, the contributions to the overall embodied carbon of the materials given by building envelope (skin), interior building components (space plan), and structure, are computed. as expected, sb #1 included a quantity of materials larger than that of sb #2. in fact, the existing materials of sb #2 are not included in the assessment. a significant volume of reinforced concrete has been used in sb #1 because only concrete provides the strength and durability performances required for some structural elements such as foundations, retaining walls, ground slabs and staircase envelope walls, whilst the structure is mainly made of wood. 3. results the histograms of figure 2 depict the results in terms of co2 eq., calculated by multiplying the quantities of materials by the unit emissions reported in the corresponding epds. in figure 2(a) the total amounts of co2 eq. of two buildings are reported and divided into the percentages of macro-layers. sb #1 accounts for the largest embodied carbon of materials, as ghg emissions are 40% higher than in sb #2. this is due to the greater amount of materials used in new constructions. on the other hand, as shown in figure 2(b) and figure 2(c), this difference shrinks when the embodied carbon is referred to the unit of gross floor area and to the number of students, respectively. the incidence of the materials of the building envelope is higher in sb #1 than in sb #2, whereas materials belonging to space plan and structure have a greater impact in sb #2. indeed, the envelope of both the buildings, including expanded polystyrene, glass and aluminium of the windows, etc., which generally have a high environmental impact. on the other hand, the structure and the internal partitions of sb #1, made of wood (which is a biogenic source of carbon storage and highly recyclable [28]), have negative values of co2 emissions (modules a1-a3 and d) and compensates the co2eq. of the high-impact materials such as concrete, steel etc. in the histograms reported in figure 3a, the three types of energy needs (heating, cooling and dhw) are compared. their incidence as a percentage of the total required energy, computed over one year of use of the two buildings, can also be observed. if on the one hand the heating requirement is the same in both the school (i.e., 76% of the required energy), on the other hand the energy for summer cooling is higher in sb #1 than in sb #2. in fact, the lower thermal inertia of wood has a marginal effect. due to the large number of labs and sport halls, hot water consumption in sb #2 is larger than in sb #1. however, the most interesting result is the higher overall energy consumption of sb #2, which is about 6% higher than that of sb #1. this percentage rises to 50% and 57% when related to the gross building area (figure 3b) and the number of students (figure 3c), respectively. such result is due to the intrinsic difficulties of making an existing building as efficient as new constructions, because some technical solutions cannot be always put into practice. for instance, in sb #2, some thermal bridges cannot be removed, and it is not possible to modify the orientation of the glazing or to install shielding systems to exploit solar radiation optimally. multiplying the energy needs by the proper emission factors, the co2 eq. emissions produced by energy requirements over the 50-year lifespan of the building can be calculated. figure 4a summarizes the overall emissions, and the related percentages, due to both materials and energy systems. however, these results are strictly dependent on the assumptions made and boundary conditions, and therefore cannot be generalised. furthermore: • as shown, global co2 eq. emissions are strongly affected by winter heating demand (both sb#1 and sb#2 are located in an alpine area). therefore, different results can be obtained if the same analysis are performed in a warmer climate scenario, where the effects of thermal inertia on the thermal behaviour of the building prevails over the transmittance of the building envelope. • with respect to the embodied carbon of materials, reference has been made to the primary data contained in epds. therefore, the result is strongly influenced by the scenarios assumed by the producers, in particular related to the end of life of the material. for example, the company producing the wood elements assumed their complete recycling at the end of life. • a lifespan of 50 years has been assumed for all building components in both schools. actually, they usually differ according to their function [25– 27], yet it is difficult to reliably estimate their lifespan since it depends on the intrinsic characteristics of the material and the local practices. besides, emission factors, used to convert operational energy into co2 eq. emissions, are assumed to be constant over the lifespan, yet increasing the use of renewable energy sources and improving efficiency of energy production plants is expected in the future, which might lead to a decrease in the emissions related to energy needs. • the environmental impact of the excavation of foundations in sb#1, along with materials demolished in sb #2 and sent to landfill or to the recycling, renovation, reuse chain was not considered due to lack of information. however, it is reasonable to assume that their influence on the total co2eq. computation is small. also, the consumption of undeveloped land, due to the construction of the new school on a vacant lot, is not taken into account. 354 vol. 33/2022 environmental performances of new and renovated school buildings 4. conclusions and limits of the study based on the above results, the following conclusions can be drawn: • the total embodied carbon of the materials of sb #1 is higher than in sb #2, both in absolute value and also related to gross floor area and number of students. this is due to the larger amount of materials used in the new building, although a significant quantity of co2 eq. is compensated by the negative values of the timber structures (wood is a biogenic source of carbon storage). • as far as energy requirements are concerned, sb #2 has higher needs. even though energy efficiency improvements are implemented on sb #2, it is not possible to reach the standards of a new building, especially with regard to the reduction in heating energy needs. however, the same analysis performed in warmer climates, where energy demand mainly relies on the thermal inertia rather than the transmittance of the envelope, gives opposite results. • the overall emissions of the two buildings are very similar. nevertheless, as sb #1 has a larger area and more students, it can be stated that sb #1 performs better than sb #2 from an environmental point of view. in fact, emissions calculated over the lifespan of 50 years override the embodied carbon of the building materials and penalize the global performances of sb #2. this study shows that, in a specific climatic context, characterized by harsh winters and mild summers, the solution of building a new school building with eco-friendly materials, such as wood, should be preferred to the refurbishment of an existing school in reinforced concrete. thus, the outcomes of the research can be implemented into a decision-making process that compare benefits and drawbacks from the two possible alternatives, within the public building sector. references [1] c. lombardi, m. baldi. cinquant’anni di economia marittima in italia : evoluzione e prospettive tra xx e xxi secolo : 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[28] g. guest, r. m. bright, f. cherubini, et al. consistent quantification of climate impacts due to biogenic carbon storage across a range of bio-product systems. environmental impact assessment review 43:21-30, 2013. https://doi.org/10.1016/j.eiar.2013.05.002. 356 https://doi.org/10.1126/science.213.4511.957 https://doi.org/10.1007/s11367-015-0938-z http://oz.berkeley.edu/users/breiman/using_random_forests_v4.0.pdf https://doi.org/10.1016/j.scitotenv.2016.03.098 https://doi.org/10.1088/1755-1315/225/1/012019 https://doi.org/10.1016/j.eiar.2013.05.002 acta polytechnica ctu proceedings doi:10.14311/app.2020.26.0013 acta polytechnica ctu proceedings 26:13–18, 2020 © czech technical university in prague, 2020 available online at https://ojs.cvut.cz/ojs/index.php/app size effect on the ultimate drying shrinkage of concrete experimental evidence and engineering practice lenka dohnalová∗, petr havlásek czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: lenka.dohnalova@fsv.cvut.cz abstract. the design and behavior of creep-sensitive structures can be to a large extent dependent on the evolution of drying shrinkage. drying shrinkage of concrete measured on standard laboratory specimens is very time demanding, furthermore, the extrapolation based on the short-term measurements is an ill-posed problem and can lead to large errors. additionally, the final magnitude of shrinkage is size-dependent which makes the transition from the laboratory data to real structure even more challenging. the present paper evaluates this size effect based on the data gathered in the creep and shrinkage database developed at the northwestern university and the data from the literature. finally, this size effect is compared against the current codes of practice and recommendations. keywords: drying shrinkage, size effect, concrete, design code, prediction model. 1. introduction concrete drying is followed by its gradual contractive volume changes referred to as drying shrinkage. similarly to autogenous shrinkage, which happens even under sealed conditions and is characteristic for modern high-performance concretes with small water-to-cement ratios, also the drying shrinkage is considered to be bounded. however, on contrary to autogenous shrinkage which is believed to be uniform within the volume of the member, drying shrinkage which is driven by the changing field of relative humidity gives rise to non-uniform stresses and can result into surface cracking. the drying shrinkage of concrete is a complex phenomenon which is not fully understood even nowadays. the evolution and magnitude of drying shrinkage is influenced by many factors, and for this reason the codes of practice and the design codes comprise many parameters. the aim of this paper is to assess the influence of the size of concrete member on the final magnitude of the drying shrinkage. in the case of concrete structures sensitive to creep and shrinkage, it is recommended to perform shortterm measurements on laboratory specimens and consequently, if necessary, to adjust the parameters of the computational models based on the results. the purpose of these short-term measurements is to capture the influence of parameters such as the water-tocement ratio, cement type, type of aggregate, additives, admixtures, etc., which significantly affect the behavior of concrete. the updating of concrete compliance function works reasonably well. in the case of drying shrinkage the final value cannot be reached in a reasonable time; not even in standard tests on laboratory specimens, 75 × 75 × 285 mm according to astm c157 [1] and 100 × 100 × 400 mm as given by the iso 1920-8 [2] standard. in both cases the values stabilize after approximately 1,000 days of drying [3] which is unacceptable in civil engineering practice. typically, the already measured drying shrinkage evolution can be accurately captured with a very simple formula with two parameters only: the characteristic time and the ultimate shrinkage. despite that the drying shrinkage updating remains an ill-posed problem [4]. the estimated ultimate shrinkage strongly depends on the choice of the characteristic time which is unknown. both drying shrinkage magnitude and the characteristic time are functions of concrete composition, dimensions and ambient humidity which influence concrete diffusivity and thus the drying rate. if the size effect on drying shrinkage magnitude is fully understood and described, then the problem might be overcome by decreasing the specimen size, the characteristic time is approximately a square function of the characteristic size. the results obtained on smaller specimens might be polluted by the loss of representativeness if the ratio of the specimen size to the aggregate is too small. however, this influence can be incorporated in the size effect law if the governing mechanisms are fully understood. the size of typical structural members exceeds the characteristic dimension of the laboratory specimens many times. for this reason, the duration of drying and shrinkage can easily reach decades until these processes cease. the currently available size effect laws on the ultimate drying shrinkage are based only on relatively limited laboratory measurements, beyond that range their applicability is questionable. two recent studies have already approached this topic. the first one [5] compares the size effect on 13 https://doi.org/10.14311/app.2020.26.0013 https://ojs.cvut.cz/ojs/index.php/app lenka dohnalová, petr havlásek acta polytechnica ctu proceedings drying shrinkage according to ec2 with two experimental series only and concludes that even the current coupled fem simulations are unable to capture the phenomenon. the second, more recent paper [6] is oriented more on the extrapolation of the drying shrinkage rather than on the size effect law, however, the original data from this paper are adapted in the present study. the first part of this paper summarizes the size effect on the ultimate value of drying shrinkage as it is described in the codes of practice and recommendations. the following section examines the relevant experimental data some of which served, very probably, as a reference used for the development and calibration of the codes and prediction models. the final part compares the models with the experimental data and summarizes the conclusions. 2. design codes and prediction models this section summarizes the definitions of the size effect on drying shrinkage according to the most common codes of practice — the american standard aci 209.2r–08 [7] and eurocode 2 [8] — and selected recommendations for the long–term behavior of concrete — b3 model [9], b4 model [10] and fib model code 2010 [11]. in the models the total shrinkage is either taken equal to the drying shrinkage (ec2, aci 209, b3) or is split into drying and autogenous shrinkage (fib mc 2010, b4), both being bounded. the influence of specimen size on time evolution of drying shrinkage is not of interest in the present study and for this reason the related formulae are not presented here. 2.1. models b3 and b4 even though the newer b4 model introduces the split of total shrinkage into drying and autogenous shrinkage, the general structure of the formula for drying shrinkage is adapted from its predecessor, model b3 whose capability is limited only to concretes with higher water-to-cement ration and which recognizes only drying shrinkage. the b3/4 models reflect not only the influence of the specimen size, but also its shape. these parameters play role both in the evolution of shrinkage and in the formula for the ultimate shrinkage. different shapes of the specimens can be treated by means of the cross-section shape factor ks (1.0–1.55) which enters the formula for the shrinkage halftime and which influences the magnitude of the ultimate shrinkage. the specimen size is expressed using the effective cross-section thickness d defined as d = 2 v/s (1) where v and s are the volume and drying surface of the member, respectively. 0.4 0.6 0.8 1 1.2 1.4 0 100 200 300 400 500 600 s iz e ef fe ct effective thickness, d [mm] b3, t0 = 7 b3, t0 = 28 ec2 aci factor aci coefficient figure 1. size effect on ultimate drying shrinkage evaluated according to selected design codes and prediction models. the ultimate shrinkage is defined by equation ε∞sh = ε ∞ s e(607) e(t0 + τsh) (2) where ε∞s is a size independent constant depending on concrete composition, curing method and 28-day compressive strength, t0 is the age at the onset of drying, e is the modulus of elasticity and τsh is a shrinkage half-time expressed as τsh = kt (ks · d)2 (3) where kt is parameter depending on 28-day compressive strength. red color in fig. 1 corresponds to the size effect on drying shrinkage magnitude given by the b3 model. from the figure it is evident that the size effect is observable only in the case of very small specimens while for structural members with realistic dimensions it completely vanishes. the size effect on drying shrinkage indirectly changes with the parameters which are in a certain way related to the evolution of stiffness during the initial phase of drying, especially the onset of drying t0 and the compressive strength fc. however, these parameters only slightly influence the shape of the curve for very small specimens, d < 50 mm, and for d > 100 mm the size effect remains negligible. 2.2. eurocode 2 similarly to the b3 model also the eurocode 2 recognizes the influence of the specimen size on the drying shrinkage rate and on the ultimate value. the specimen shape does not come into the calculation and the specimen size is represented by the notional size defined as h0 = 2 ac/u (4) where ac is the area of cross-section and u is the perimeter of the member in contact with ambient environment. this expression coincides with (1) in the case of very (infinitely) long prismatic members. 14 vol. 26/2020 example of an article with a long title the ultimate shrinkage strain is adjusted by factor kh which is a decreasing piece-wise linear function of the notional size. the reference value kh = 1 is specified for h0 = 100 mm, kh = 0.85 for h0 = 200 mm, kh = 0.75 for h0 = 300 mm and kh = 0.7 if h0 reaches or exceeds 500 mm. for smaller notional sizes than 100 mm the code does not provide any guidance. the size effect is shown in fig. 1 by blue dashed line. 2.3. aci 209.2r–08 the size effect on drying shrinkage is in the american standard introduced by means of multiplicative correction terms which affect the ultimate value of shrinkage strain. the evolution of drying shrinkage strain in time can be according to this standard treated by two distinctly different ways. similarly to other standards, the first method prescribes a time function which depends on the v/s ratio; on contrary to this, the second method enables to use constant parameters independently of the size and shape of the concrete member and thus this method gives size-independent results. the correction term reflecting the size effect on the ultimate drying shrinkage is introduced by two different ways, either it is called a coefficient or a factor. in order to be consistent with the preceding sections and the figures, here we slightly modify the original formulae presented in the code, because the aci uses a different definition of the characteristic thickness. the size-dependent correction coefficient is prescribed by a single expression γsh,vs = 1.2 exp (−0.00236 d) (5) the correction factor uses a piece-wise linear decreasing function for members with 25 mm < d < 75 mm, for which the value is in the range 1.35–1.00, for larger members 75 mm < d < 190 mm it offers a simple linear expression γsh,d = 1.17 − 0.00228 d (6) the correction coefficient and factor are shown in fig. 1 by solid black and dash-dotted black lines, respectively. 2.4. fib model code 2010 identically to ec2, the specimen size is in fib mc 2010 expressed by means of the notional size given by (4). yet, the notional size affects only the shrinkage evolution which becomes faster with decreasing size. the ultimate shrinkage is entirely independent of the specimen size, for this reason the size effect does not need to be illustrated in fig. 1. 3. experimental data the experimental studies in which the evolution of drying shrinkage was measured on specimens with different sizes allow for identification of the size effect 0 100 200 300 400 10 100 1000 s h ri n k ag e, ε sh [ 1 0 -6 ] drying duration, t-t0 [days] d = 31 mm d = 62 mm d = 93 mm d = 311 mm figure 2. the experiment of l’hermite (data adapted from [16]) shows incomplete drying shrinkage of the largest specimen with d = 311 mm. such measurements were not discarded but are denoted by empty symbols in fig. 4. law on drying shrinkage. naturally, the specimens had to be manufactured from the same concrete mixture and stored at the same conditions for sufficiently long period of time to reach the ultimate shrinkage strain. except for the findings reported by samouh [6], keeton [14], and bryant [12], the experimental data investigated in this study are adapted from the creep and shrinkage database developed at the northwestern university [16]. keeton’s original data on total shrinkage was preferred to the database which contains somehow modified values. the database contains 61,512 values measured on 3,308 specimens (1,439 for creep and 1,869 for shrinkage) investigated within 340 experimental surveys (143 creep, 197 shrinkage). for the purpose of the present analysis it was necessary to narrow the shrinkage database only to those studies which examined shrinkage of concrete specimens of (at least two) different sizes made of the same concrete mixture. furthermore, the curing type and duration had to be the same and so had to be the ambient relative humidity. in total 21 experimental surveys met these criteria and from these 14 had to be discarded for some reason, most often because the duration of the experiment did not suffice to obtain stable value of the shrinkage strain of at least two specimens with different size. the setup and basic material properties of the suitable experiments are summarized in table 1. 4. discussion a comparison of the size effect on drying shrinkage is presented in fig. 1. it can be immediately noticed that the european and american standards use different member size as a reference. if both curves are normalized with respect to d = 60 mm, the differences (at least for 100 mm < d < 200 mm) considerably diminish as documented in fig. 5. the eurocode 2 15 lenka dohnalová, petr havlásek acta polytechnica ctu proceedings experimental year shape, d t0 a/c c w/c fc rh tmax study [mm] [day] [-] [kg m−3] [-] [mpa] [%] [day] bryant p [12] 1987 p 50–200 8 5.09 390 0.47 50.0 60 1,950 bryant s [12] 1987 s 100–400 8 5.09 390 0.47 50.0 60 1,880 hansen † [13] 1966 c 46–252 8 6.00 303 0.71 41.3 50 1,750 keeton 20 [14] 1965 c 33/44/65 8 3.74 452 0.46 46.5 20 897 keeton 50 [14] 1965 c 33/44/65 8 3.74 452 0.46 47.5 50 897 keeton 75 [14] 1965 c 33/44/65 8 3.74 452 0.46 46.8 75 897 l’hermite † [15] 1970 p 31–311 1 5.06 350 0.51 na 55 699 mazloom † 2004 c 35/60 7 3.7–4.4 425–500 0.35–0.41 58–70 50 586 samouh i [6] 2017 c 39/57/82 1 4.89∗) 320∗) 0.59∗) 49.6 50 300 samouh ii [6] 2017 c 39/57/82 1 6.03∗) 258∗) 0.74∗) 32.6 50 300 shritharan † 1989 p 44–178 8 5.09 393 0.47 50.1 60 3,492 wittmann † 1991 c 33/64/120 7 5.43 350 0.48 33.2 65 1,105 table 1. details of the experimental setup of the selected shrinkage experiments. the specimen shape is either a prism (p), cylinder (c), or a slab (s), d is the effective thickness, t0 is the onset of drying, c, w, a is the cement, water and aggregates content, respectively, in the concrete mixture. the average compressive strength on cylinders at 28 days is denoted f c, rh is the ambient relative humidity, and tmax is the duration of the experiment (i.e. drying duration). in samouh i and ii, asterisks ∗) indicate that c corresponds to the content of opc, not to the total amount of cementitious material, the data with symbol † are adapted from the nu database [16]. 0 200 400 600 800 0.1 1 10 100 1000 s h ri n k ag e, ε sh [ 1 0 -6 ] drying duration, t-t0 [days] d = 44 mm d = 67 mm d = 89 mm d = 133 mm d = 178 mm sealed figure 3. evolution of drying shrinkage measured by shritharan, the data adapted from [16] serve as an example of experiment in which the shrinkage of all specimens can be considered to have terminated. does not define size effect on drying shrinkage magnitude for d < 100 mm; the value for 60 mm was linearly extrapolated from the interval 100–200 mm. in contrast to the other models, the fib mc 2010 and for d > 100 mm also the b3 model completely neglect the size effect on the ultimate shrinkage. compared to other models, the b3 model is the only model which incorporates the onset of drying into the size effect on the ultimate drying shrinkage. however, this effect is limited only to specimens with d < 30 mm which makes the purpose of the entire proposed and rather complex formula (2) rather questionable. a comparison of the processed experimental data is displayed in fig. 4. the value of shrinkage strain at the end of the experiment was normalized with respect 0.4 0.6 0.8 1 1.2 1.4 60 0 100 200 300 400 n o rm al iz ed u lt im at e sh ri n k ag e, ε~ s h ,∞ effective thickness, d [mm] wittmann shritharan l′hermite bryant s bryant p mazloom keeton 20 keeton 50 keeton 75 hansen samouh i samouh ii figure 4. influence of the effective thickness on the normalized experimentally measured ultimate shrinkage strain. the strain is normalized with respect to the value corresponding to d = 60 mm, empty symbols mark the last data points in prematurely terminated experiments. to the value corresponding to d = 60 mm (if missing, the value was obtained by the linear interpolation or linear extrapolation such as in the case of “slabs” specimens from bryant’s experiment). the filled points denote the experiments which were treated as terminal while the empty marks correspond to the experiments which did not last sufficiently long and in which further increase of shrinkage can be expected. the symbol type was selected to resemble the shape of the specimen (circle for a cylinder, square for a prism). as documented by the fig. 4, the experiments show a uniform and almost linearly decreasing trend in the range 30 mm < d < 150 mm. however, there are almost no data for d > 150 mm. this does not infer 16 vol. 26/2020 example of an article with a long title that the experiments above 150 mm do not exist at all but rather the experimental setup did not explore different sizes. a very small size effect on drying shrinkage can be observed in the case of the “slab” specimens tested by bryant. however, the results presented in fig. 4 might be misleading: the smallest specimen size was d = 100 mm while the values are normalized with respect to d = 60 mm. this was done by extrapolation from the experiment with d = 100 mm and d = 150 mm which had almost the same value of shrinkage. similar result offers also l’hermite in whose experiment the specimens started drying already at the age of 1 day but very probably this is not the only reason. on the other hand, the steepest size effect on drying shrinkage was documented by keeton. his experiment also demonstrates that the size effect is humidity independent. yet, there is no experimental evidence which could provide a clue to identification of the size effect in more massive members. it must be noted that the ultimate drying shrinkage of thick members might not be of interest because the design life-time might actually come earlier than the member dries. figure 5 compares the experimental data with the prediction models. the models which do not consider size effect on drying shrinkage — fib mc 2010 and model b3 (above 100 mm) — are evidently incorrect. the best agreement is reached with both alternative formulations in aci 209 which somehow create a lower and upper-bound to almost all experimental data. since the eurocode 2 provides a recommendation only for d > 100 mm where very few experimental data are present, it is difficult to assess its performance. however, the slope of the ec2 curve in fig. 5 is similar to the aci 209 which provides a good agreement. unfortunately, the newly prepared version of eurocode 2 [17] which origins from the model code 2010 omits the size-dependent coefficient kh from the formulation and thereby completely neglects the size effect on the ultimate drying shrinkage. this will lead to more conservative but less economic design. 5. conclusions the presented study analyzed the size effect on the ultimate drying shrinkage of concrete and compared the processed experimental data to the recommendations and codes of practice. the following conclusions can be drawn: • according to the currently valid codes of practice, namely the eurocode 2 and aci 209.2r-08, the ultimate drying shrinkage is decreasing with increasing size (effective thickness). it can be stated that for members with effective thickness d between 100 mm and 200 mm this size effect is captured similarly. 0.4 0.6 0.8 1 1.2 1.4 60 0 100 200 300 400 s iz e ef fe ct effective thickness, d [mm] b3 ec2 aci factor aci coefficient figure 5. comparison of the size effect according to the design codes and recommendations (curves) with the experimental data (filled black symbols = finished shrinkage, empty grey symbols = prematurely terminated experiments). the functions and the data are normalized with respect to the size d = 60 mm. • as described by the b3/4 models the size effect is a function of the age when concrete starts drying and is very pronounced only for specimens with d < 50 mm, otherwise the size effect almost vanishes, similarly to the fib model code 2010 which ignores this size effect completely (and very probably so will do the new ec2). • the experimental database [16] contains 6 suitable experimental studies which together with 2 additional studies from the literature provide consistent results confirming the presence of size effect on the ultimate drying shrinkage. none of the experimental studies exhibits opposite or no size effect on the drying shrinkage magnitude. however, it must be noted that sound experimental data for d > 150 mm are missing and thus cannot provide any evidence above this limit. • for d in the range from 20 mm to 150 mm both approaches of the aci 209 code exhibit by far the best agreement with the experimental data. • the experimental data on the ultimate shrinkage of specimens with d > 200 mm are not available and very likely will not be available in the near future. in opinion of the present authors, the code-like expressions should not be constructed by a simple extrapolation beyond the experimentally explored range. more preferably, these expressions should origin from the fem simulations exploiting properly calibrated and physically based material model which will perform well in the well-documented range 20 mm < d < 150 mm. acknowledgements the authors gratefully acknowledge financial support from the grant agency of the czech technical university in prague, project number sgs19/032/ohk1/1t/11, and from the czech science foundation, project number 19-20666s. 17 lenka dohnalová, petr havlásek acta polytechnica ctu proceedings references [1] american society of testing and materials. astm c157, standard test method for length change of hardened hydraulic-cement mortar and concrete. 2008. [2] international organization for standardization. iso 1920-8 testing of concrete part 8: determination of drying shrinkage of concrete for samples prepared in the field or in the laboratory. 2009. [3] c. aguilar. estudio del comportamiento y desarrollo de una metodologia de prediccion de la retraccion hidraulica de hormigones, phd. thesis (in spanish). pontificia universidad catolica de chile. escuela de ingenieria, 2005. [4] z. bažant, m. jirásek. creep and hygrothermal effects in concrete structures. springer netherlands, 2018. doi:10.1007/978-94-024-1138-6. [5] j. m. torrenti, f. benboudjema. desiccation shrinkage of large structures: is there a size effect? pp. 404–411. 2013. doi:10.1061/9780784413111.048. [6] h. samouh, e. rozière, a. loukili. influence of specimen size and measurements duration on the long-term extrapolation of drying shrinkage. construction and building materials 150:276 – 286, 2017. doi:10.1016/j.conbuildmat.2017.06.009. [7] aci committee 209. guide for modeling and calculating shrinkage and creep in hardened concrete. aci, 2008. [8] european committee for standardization. en 1992-1-1 eurocode 2: design of concrete structures–part 1-1: general rules and rules for buildings. cen, brussels, 2005. [9] z. bažant, s. baweja. creep and shrinkage prediction model for analysis and design of concrete structures: model b3. adam neville symposium: creep and shrinkage structural design effects 2000. [10] rilem technical committee tc-242-mdc (z.p. bažant, chair). model b4 for creep, drying shrinkage and autogenous shrinkage of normal and high-strength concretes with multi-decade applicability (rilem technical committee tc-242-mdc multi-decade creep and shrinkage of concrete: material model and structural analysis). materials and structures 48:753–770, 2015. [11] fédération internationale du béton. model code 2010. no. vol. 65 in fib bulletin. international federation for structural concrete (fib), 2012. [12] a. h. bryant, c. vadhanavikkit. creep, shrinkage-size, and age at loading effects. aci materials journal 84:117–123, 1987. [13] t. hansen, a. mattock. influence of size and shape of member on the shrinkage and creep of concrete. journal of the american concrete institute 63-10:267–290, 1966. [14] keeton. study of creep in concrete, phase 1 (i-beam). u. s. naval civil engineering laboratory, port hueneme, california, 1965. [15] r. l’hermite, m. mamillan, c. lefevre. nouveaux résultats de recherches sur la déformation et la rupture du béton. annales de l’institut technique du bâtiment et des travaux publics 18(207-208):323–360, 1965. [16] m. hubler, r. wendner, z. bažant. comprehensive database for concrete creep and shrinkage: analysis and recommendations for testing and recording. aci 112:547–558, 2015. [17] european committee for standardization. eurocode 2: design of concrete structures–part 1-1: general rules, rules for buildings, bridges and civil engineering structures, final version of pt1-draft pren 1992-1-1 2018 d3. 2018. 18 https://doi.org/10.1007/978-94-024-1138-6 https://doi.org/10.1061/9780784413111.048 https://doi.org/10.1016/j.conbuildmat.2017.06.009 acta polytechnica ctu proceedings 26:1–6, 2020 1 introduction 2 design codes and prediction models 2.1 models b3 and b4 2.2 eurocode 2 2.3 aci 209.2r–08 2.4 fib model code 2010 3 experimental data 4 discussion 5 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0078 acta polytechnica ctu proceedings 33:78–84, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague predicting leakage of the vercors mock-up and concrete containment buildings a digital twin approach laurent charpina, ∗, jessica haelewynb, anass cherki el idrissib, julien niepceronc, benoît massonc, charles toulemondea, guillaume boulantd, jean-philippe mathieua, françois hamonb, sylvie michel-ponnelleb, jean-marie hénaulte, frédéric tailladee, jean-luc adiaa, florian escoffierb a edf r&d mmc, les renardières, avenue des renardières, 77250 ecuelles, france b edf r&d ermes, saclay, 7 boulevard gaspard monge, 91120 palaiseau, france c edf dt, 19 rue pierre bourdeix, 69000 lyon, france d edf r&d pericles, saclay, 7 boulevard gaspard monge, 91120 palaiseau, france e edf r&d prisme, chatou, 6 quai watier, 78400 chatou, france ∗ corresponding author: laurent.charpin@edf.fr abstract. edf operates a nuclear power generation fleet made up of 56 reactors. this fleet contains 24 reactors designed as double-walled concrete containment building. the inner concrete containment vessel has no metallic liner and is a prestressed reinforced concrete building. the inner concrete containment vessel is designed to withstand a severe accident, in terms of mechanical and sealing behaviour. the tightness of the containment is tested every 10 years, by carrying out a pressurization test and by measuring the leak rate. the leak rate is required to be below a regulatory threshold to continue operation of the concrete containment building for the next ten years. ageing of concrete due to drying, creep and shrinkage leads to increase prestress loss and then leak rate with time. for some containment buildings, the leak rate gets closer to the regulatory threshold with time, so important coating programs are planned to mitigate and limit the leak rate under the regulatory threshold. therefore, it is very important for edf to have a concrete containment building leak rate prediction tool. to address this issue, an important research program around a 1/3 scale concrete containment building mock-up called "vercors" have been launched at edf. the mock-up is heavily instrumented, and its materials (concrete, prestressing cables) have been widely characterized and studied. an important numerical effort has also been made to implement structural computations of the mock-up and to capitalize these computations as well as their post-processing (so as to compare automatically with the monitoring data) in what can be called a digital twin of the mock-up. this digital twin is now used to predict the leakage of vercors mock-up before yearly pressure test, and also to optimize the repair programs on the real containments. keywords: air leakage, concrete containment building, creep, drying, shrinkage. 1. introduction 1.1. industrial context edf is the main electricity utility in france and operates a 56 reactors fleet of nuclear power plants (npps). pwr safety relies on three confinement barriers: fuel cladding, primary circuitry integrity, and reactor containment building. concrete containment buildings (ccbs) leak-tightness is checked every 10ăyears during the integrated leakage rate test (ilrt). this test must be performed and validated in order to continue the plant operation [1]. double-wall containment buildings leak-tightness is achieved thanks to the inner prestressed reinforced ccb (without a metallic liner) and dynamic air filtering between the two containment buildings. in case of an accident, the concrete remains in compression. the design of the containment building is calibrated to this aim. in france, the choice has been made to use grouted tendons, therefore increasing posttension is not an option (which is singular compared to ccbs in other countries). edf and its industrial partners has acquired a vast experience of repair solutions that are thoroughly tested and qualified in the lab and have shown good efficiency in the field. however, their operational implementation remains difficult due to accessibility and schedule considerations, and therefore needs to be optimized. to perform this optimization and ensure a high level of safety of the containments, knowledge about concrete behaviour and capability to predict concrete delayed strains as 78 https://doi.org/10.14311/app.2022.33.0078 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 predicting leakage of the vercors mock-up and ccbs well as leak-tightness evolution are required. the vercors programme is the current main effort at edf to bridge the gap between material scale knowledge of concrete behaviour and leak-tightness prediction. 1.2. vercors mock-up, a unique experiment a 1/3 scale mock-up of a double-wall ccb has been built by edf in the south of paris (edf lab les renardières). it is called vercors (vérification réaliste du confinement des réacteurs / realistic assessment of reactors containment, figure 1). for more details on the project and the first international simulation benchmark (organized in 2015 on vercors early-age behaviour), the reader can refer to [2]. the second vercors benchmark about long-term mechanical behaviour and leak-tightness has been reported in [3]. a third benchmark has started in january 2021. in addition to cost, the huge advantage of the scale reduction is the acceleration of drying: since the mock-up is 1/3 compared to real containment buildings, it is expected to dry 9 times faster. in addition, the 1/3 scale makes it possible to reduce all featured of the mock-up (tendons, rebars, etc.), therefore representativity to real containment buildings is very satisfactory. aggregate size and sensors have not been reduced by the same ratio for practical reasons. vercors was built starting in 2013 in less than 1 year. prestressing started after completing concrete casting, at least 28 days after casting of the dome. the first irlt has been performed in the end of year 2015. in march 2016 the air conditioning system, ensuring representative temperature and humidity conditions both in the inner containment and between the two containments, has been started. then, irlts have been performed every year (except during the covid lock-out period), to represent the tests performed every 10 years on real ccbs. instrumentation has been the focus of large efforts, with both the classical instrumentation found on ccbs but in much larger numbers (more than 300 vibrating wire strain sensors, 200 pt100 temperature sensors, pendula and invar wires) and also innovative sensors (2 km of optic fibers, 20 tdr moisture sensors, 80 strain gages on rebars). vercors concrete has been thoroughly characterized in edf and national and european partner labs to achieve a consistent knowledge of elastic properties, strength, drying properties, delayed strains at different ages and different temperatures. more information can be found in [4]. 1.3. leak-tightness model strategy the strategy to use vercors to be able to predict the ccb behaviour during irlts and to optimize coating programs relies on the construction of a ccb digital twin. it is based on physical ideas described in [5]. the present paper is the focus of the two first steps, as described on figure 2: • validation of the physical models (for drying, shrinkage and creep at moderate temperature) using laboratory data obtained in edf laboratories and in macena pia project and a comparison of the staggered thermo-hygro-mechanical analysis to vercors results. • calibration of a global analytical leakage model on existing vercors ilrt results, which is then used to predict the leakage during next irlts. 1.4. objective of the present paper in this paper, the procedure presented in figure 2 to the vercors mock-up is followed. the main assumptions of the structural computation will be presented as well as the results obtained. then, the global leakage model and its results will be presented. 2. vercors structural computation the objective of this t-h-m computation is twofold: validating the models involved by comparison to the monitoring data on the vercors mock-up, and providing inputs to the leak-tightness model. 2.1. model assumptions most assumptions of the vercors modelling have been presented elsewhere [2], so only the main features as well as references will be provided in the present paper. all computations are performed with code_aster (www.code-aster.org) in the framework of the vercors digital twin developed at edf. 2.1.1. constitutive modelling the thermal model is a linear conductivity model. the thermal model parameters (heat capacity, conduction coefficient) have been measured on vercors concrete. the drying model is that developed in granger’s works from mensi’s earlier developments [6]. it is a non-linear diffusion model using water concentration as an unknown (hence, boundary conditions needs to be expressed in water concentration which requires the knowledge of the desorption isotherm, since the ambient moisture is known from air relative humidity). the drying model and calibration issues are presented in detail in [6], and the parameters used in this study are available in [7]. the mechanical model is based on benboudjema and sellier’s works as well as internal edf works [8–10]. it is an additive model taking into account elasticity, drying shrinkage, basic creep and drying creep. the calibration of the model is described in [7] as well. this calibration phase involving a large number of drying, shrinkage and creep experiments at 20 ◦c and 79 www.code-aster.org l. charpin, j. haelewyn, a. cherki el idrissi et al. acta polytechnica ctu proceedings ������$�wklug�ehqfkpdun�zloo�vwduw�rq�-dqxdu\����������� � figure 1. outer containment wall of the vercors mock-up and sectional schematic view figure 2. model validation strategy proposed in the present paper, corresponding to points 1 and 2 of the above strategy. 40 ◦c yields a set of parameters which is used in the present study. concerning the properties of prestress cables, they were communicated by fressinet who did the posttensioning. the relaxation of prestressing cables is not accounted for. passive reinforcements are not taken into account in the model. the mechanical model is based on benboudjema and sellier’s works as well as internal edf works [6, 7, 11]. it is an additive model taking into account elasticity, drying shrinkage, basic creep and drying creep. the calibration of the model is described in [5] as well. this calibration phase involving a large number of drying, shrinkage and creep experiments at 20 ◦c and 40 ◦c yields a set of parameters which is used in the present study. concerning the properties of prestress cables, they were communicated by fressinet who did the posttensioning. the relaxation of prestressing cables is not accounted for. passive reinforcements are not taken into account in the model. 2.1.2. fe model the numerical model of the vercors mock-up is focused on the inner pre-stressed concrete containment building. the mesh used for thermal and hydric computations is composed of 331962 nodes and 304928 hexahedral linear elements, while the mesh used for the mechanics computation is composed of 211776 nodes and 40656 hexahedral quadratic elements (figure 3). 80 vol. 33/2022 predicting leakage of the vercors mock-up and ccbs � figure 3. vercors mesh. the pre-stress cables are individually modelled using bar elements. deviations around the access hatch and other passages through the walls are represented. the metal elements closing the through-wall passages (pipes, hatches) are also modelled in 3d. the mesh is refined at the surface of concrete to correctly capture the drying gradients. in the cylindrical part, the wall thickness is discretized using 12 elements which size vary between 2 cm and 6 cm. 2.1.3. boundary conditions and loadings the thermal and hydric boundary conditions are very important since drying, creep and shrinkage are highly influenced by temperature and humidity. the ambient temperatures and humidity are the results of: • the external weather (especially in the early months of the experiment during which the external containment was not completed). • the behaviour of the air-conditioning system which is designed to impose temperature and humidity similar to field conditions (both inside the containment and in the space between internal and external containment), but needs to be stopped for about 3-4 weeks for ilrts. • the supply of humidity due to the flooding of the basement raft during ilrts (which is related to the procedure also used on real ccbs to measure the leakage through the containment wall only), which brings moisture close to 100 % rh during about 2 weeks. in this work, the boundary conditions used are a schematic representation of the available measurements. for the temperature, the temperature sensors installed on the wall surface are used (due to significant boundary layer effects, the wall temperature is quite different to ambient temperature, which has also been modelled directly in other works, but is not presented here). for moisture, the rh measurements performed on both sides of the containment walls are used. the mechanical boundary conditions are as follows: displacements at the bottom surface of the basement raft are assumed to be zero. the post-tension of the cables is performed in 8 phases instead of the 35 real phases. it was shown to be equivalent in terms of post-tensioning global and local mechanical behaviour. finally, and most importantly, the computation starts at the beginning of drying of concrete (shortly before the beginning of prestress). hence, early-age effects are not represented. then, the mechanical fields must be understood as relative to the mechanical state at the end of the early-age period. 2.1.4. vercors digital twin the data associated to vercors (monitoring, materials), the parameters identification procedures, the structural computation of vercors and related post-processing functions have been capitalized in what can be described as the digital-twin factory of vercors.) the aim of such a tool is to improve quality and traceability of the studies, and also to save time when changes are performed to a part of the study. the different tools needed for building such a digital twin are described in a more detailed manner in [4]. 2.2. results of the staggered t-h-m analysis modelling in this section, some comparison to monitoring results are presented, and some conclusions are drawn on the validation of the models. 81 l. charpin, j. haelewyn, a. cherki el idrissi et al. acta polytechnica ctu proceedings figure 4. simulated and measured moisture (bottom of the wall, see arrow) 2.2.1. hydric behaviour first, the results of the hydric computation are compared to moisture sensor results available on the vercors mock-up. these results have been obtained using tdr sensors [6, 12, 13]. the comparison is shown for one of the sensors on figure 4. the set of parameters studied here is the set called "2018" on figure 4. it has been calibrated on drying experiments performed on well-hydrated samples at rh close to 50 %, see [7]. the simulation assumes an initial saturation ratio of 1 which is consistent with the fact that the vercors concrete has been artificially kept wet during construction. the "2016" set of parameter performs poorly and is considered as non-representative of the field concrete since it was calibrated on a drying test performed at 24h (nonmature concrete). the 2018 model does not perfectly match the experiment: a fast drying is predicted before the heating system is turned on (600 days or march 2016), whereas for the tdr post-processing, it has been assumed that the saturation is still equal to 1 when heating is turned on. 2.2.2. vercors digital twin this inability to reproduce the moisture field with the "2018" identification lead to different investigations: • as explained in [6], the calibration of granger model on mass-loss data is an ill-posed problem. the set of parameters might be inadequate for this reason. the calibration is now being improved by using other sources of data: mass-loss at different rhs, moisture profile measurements in large samples in vercors mock-up. • the model itself might be inadequate for the extrapolation from small laboratory samples to large samples. therefore, a new model is currently tested on vercors mock-up based onărichardfick’s law [14]. figure 5. simulated and measured strain (middle part of the wall, see arrow) • the calibration procedure for the tdr sensor, linking permittivity to saturation, is also a complicated process [13], a precise temperature dependence correction is required. as a comparison, another set of parameters called "2020" has also been used. the identification of the thermal activation coefficients of the model has been refined, which enables predicting more accurately the slow drying before heating is turned on. 2.2.3. mechanical behaviour the strain computed in the mechanical computation are then compared with measured strains. it is chosen here to use the vibrating wire sensors as a reference as they are available on ccbs. as an example, the comparison is performed on one sensor (at mid-height) on figure 5. the total strain is represented (the strain is corrected from effects of temperature on the sensor, but not from thermal dilation effects of the structure). the comparison of simulation and measurements for the "2018" set of parameters is quite satisfactory. many conclusions car be drawn: • the low strain occurring before the start of the ac/heating system in march 2016 is well reproduced. • the acceleration of the strain kinetics after march 2016 is also well reproduced, showing that the thermal activation is well represented (at least globally) in the model. • the swelling related to water uptake during the pressure tests (due to lower temperature and higher moisture) is underestimated. on the contrary, the previous set of parameters "2016" performs much less well. indeed, this set of parameters was only partly calibrated on vercors concrete. hence, we see that despite an unsatisfactory representation of the moisture in the concrete wall, the mechanical prediction using the "2018" set of 82 vol. 33/2022 predicting leakage of the vercors mock-up and ccbs � figure 6. local leakage prediction method. parameters are consistent with measurements. this shows that the calibration procedure of mechanical behaviour can "compensate" an unsatisfactory hydric calibration to a certain extent. to improve our understanding of this phenomenon, new studies will be started to quantify the uncertainty of the moisture and strain predictions depending on the constitutive models calibration methodology. 3. vercors leak prediction 3.1. model assumptions the leakage model is based on the assumption that leaks through the wall can be classified into two categories: • diffuse leaks that flow mostly through the concrete porosity. hence, these leaks are insensitive to the crack opening, but mostly related to the sound concrete permeability and its evolution, which is itself driven by concrete saturation. • localized leaks (which are measured during the ilrt using leak boxes which collect and measure the flux of the air flowing through localized defects). hence, these leaks are sensitive to the mechanical state of the containment since cracks can open and close due the pressure applied during the ilrt. this effect is considered to be dependant to the evolution of the prestress. 3.1.1. porosity leaks the input for the computation of the diffuse leaks is a moisture profile computed with the t-h-m model. then, a second computation is performed to reproduce the air-flux through the wall at a given saturation profile. the model used requires the knowledge of the vercors van genuchten isotherm parameters to compute the relative permeability according to mualem’s law. the model intrinsic permeability is calibrated in order to reproduce the correct porosity leak flow in "vd2" test of 2018. the model is then used to predict the porosity leakage in "vd3" test in figure 7. leakage measurements and vd3 prediction (o = measurements, x = prediction). 2019, using the same parameters but changing the saturation profile according the t-h-m computation results. the saturation profile used corresponds to the computation labelled "2016" in the previous figures, since at the time when the study was performed, it was thought to best reproduce the tdr measurements (which have since then been corrected). 3.1.2. local leaks the evolution of local leaks is assumed to be proportional to the prestress decrease (average on the full mock-up and the 2 in-plane directions in this approach). an affine relation is found between the prestress loss and the local leakage, which is used to extrapolate the value of the local leakage to the next vd by using the predicted prestress loss, as shown on figure 6. 3.2. leak-tightness prediction results using this methodology, a prediction for the leakage drying the "vd3" pressure test is obtained, only based on information which were already available before the pressure test. the results are shown on figure 7 as well as the measured result. the leakage predictions are very close to the measurements (for local and diffuse leaks). therefore, the global leakage model is considered very satisfactory. it is believed that despite being a quite simple model, it correctly captures the leakage behaviour of the mock-up. 4. perspectives in 2020 many changes will be brought to the vercors modelling: • new efforts on the drying law (validation and calibration) will be made. • our ability to do the same predictions using a reduced set of monitoring data for calibration (as they are available on ccbs) will be tested. 83 l. charpin, j. haelewyn, a. cherki el idrissi et al. acta polytechnica ctu proceedings • works a localized version of the leakage model will be continued, in order to be able to predict better where on the mock-up leakage occurs, assess the effect of local prestress loss and take refurbishment into account in a more efficient manner. moreover, a third benchmark has started in 2021. references [1] a. simon, a. courtois. structural monitoring of prestressed concrete containments of nuclear power plants for ageing management, structural mechanical in reactor technology 21:6-11, 2011. [2] m. corbin, m. garcia. international benchmark vercors 2015 overview, synthesis and lessons learned, edf, septen, 2015. [3] l. charpin, j. niepceron, m. corbin, et al. ageing and air leakage assessment of a nuclear reactor containment mock-up: vercors 2nd benchmark. nuclear engineering and design 377, 2021. https: //doi.org/10.1016/j.nucengdes.2021.111136. [4] j.-p. mathieu, l. charpin, p. sémété, et al. temperature and humidity-driven ageing of the vercors mock-up, euro-c, austria, 2018. [5] l. charpin, c. toulemonde, j. l. adia, et al. double wall containment buildings leak-tightness prediction: strategy and application, in international rilem conference on early-age and long-term cracking in rc structures, 2021. [6] l. charpin, a. courtois, f. taillade, et al. calibration of mensi/granger constitutive law: evidences of ill-posedness and practical application to vercors concrete, tince 2018, paris-saclay, france, 2018. [7] l. charpin, j. haelewyn, j.-p. mathieu. identitication of drying, creep and shrinkage constitutive laws for concrete at 20c and 40c, application to vercors mock-up, in smss, rovinj, croatia, 2019. [8] f. benboudjema. modélisation des déformations différées du béton sous sollicitations biaxiales. application aux enceintes de confinement de bâtiments réacteurs des centrales nucléaires, doctoral dissertation, université de marne la vallée, 2002. https://doi.org/10.1051/978-2-7598-1846-4.c009. [9] a. sellier, l. buffo-lacarrière. vers une modélisation simple et unifiée du fluage propre, du retrait et du fluage en dessiccation du béton. european journal of environmental and civil engineering 13(10):1161-82, 2011. https://doi.org/10.1080/19648189.2009.9693184. [10] a. foucault, s. michel-ponnelle, e. galenne. a new creep model for npp containment behaviour prediction, sscs 2012, aix en provence, france, 2012. [11] l. granger. comportement différé du béton dans les enceintes de centrales nucléaires : analyse et modélisation, ecole nationale des ponts et chaussées, ph.d. thesis, 1996. [12] d. vautrin, f. taillade, a. courtois, et al. adaptation of a tdr probe design for the estimation of water content in concrete, tince, paris, 2016. [13] v. guihard, j. haelewyn, a. courtois, et al. relevance and means for measuring water content in concrete structures. an illustration with pwr concrete containment, tince, paris, 2018. [14] j. carette, f. soleilhet, f. benboudjema, et al. identifying the mechanisms of concrete drying: an experimental-numerical approach. construction and building materials 230, 2020. https: //doi.org/10.1016/j.conbuildmat.2019.117001. 84 https://doi.org/10.1016/j.nucengdes.2021.111136 https://doi.org/10.1051/978-2-7598-1846-4.c009 https://doi.org/10.1080/19648189.2009.9693184 https://doi.org/10.1016/j.conbuildmat.2019.117001 https://doi.org/10.14311/app.2022.33.0558 acta polytechnica ctu proceedings 33:558–563, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague concrete admixtures sustainable concrete oliver schwoon sika services ag, tueffenwies 16, 8048 zurich, switzerland correspondence: schwoon.oliver@ch.sika.com abstract. durability of concrete, an important part of the sustainability theme, was, is, and will be one of the main topics for large infrastructure projects like the gotthard base tunnel. for such important structures a service life of more than 80 years is expected. the prevailing underground conditions like water ingress, aggressive waters etc. will influence the choice of the concrete system and mix design. therefore, a focus is on the potential damage mechanisms on the concrete like chloride ingress or sulfate attack. a well-designed concrete will be durable concrete which can withstand various external attacks and achieves or even extends the designed service life of the concrete structure. based on the knowledge that the main damaging mechanisms for concrete are based on water penetration into the concrete, various measures can be taken to achieve concrete structures with significantly reduced permeability. these measure ranges from reduction of the water/binder-ratio, increasing the density of the concrete using silica fume, to the use of special admixtures in order to increase the durability of concrete. in an extensive research program, the influence of mix design components e.g. different admixture types, water/binder-ratio on hardened concrete properties like chloride resistance (chloride migration coefficient), sulfate resistance and water conductivity have been tested and analyzed. keywords: admixtures, chloride migration coefficient, mix design, sulfate resistance. 1. introduction the durability of concrete, an important part of the sustainability theme, was, is, and will be one of the main topics for large infrastructure projects like the gotthard base tunnel. for such important structures a service life of more than 80 years is expected. hardened concrete can be damaged by water containing sulfates which can occur in soil or dissolved in ground water. these sulfates reacts with the tricalcium aluminate hydrates (afm) in the cement and forms ettringite (also thaumasite under certain conditions). that leads to increases in volume and to high internal pressure in the concrete structure and therefore cracking and spalling occurs [1]. another aspect is on the protection of the reinforcement against chloride exposure or chloride ingree into the concrete. chloride ions can migrate through the concrete mainly by diffusion, interact with the passivation layer of the steel and leading subsequently to corrosion of the steel reinforcement. other ways of transportation of the chloride ions are capillary absorption as well as hydrostatic pressure [1]. the prevailing underground conditions like water ingress, aggressive waters etc. will influence the choice of the concrete system and mix design. therefore, a focus is on the potential damage mechanisms on the concrete like chloride ingress or sulfate attack. a well-designed concrete will be durable concrete which can withstand various external attacks and achieves or even extends the designed service life of the concrete structure. based on the knowledge that the main damaging mechanisms for concrete are based on water penetration into the concrete, various measures can be taken to achieve concrete structures with significantly reduced permeability. these measure ranges from reduction of the water/binderratio, increasing the density of the concrete using silica fume, to the use of special admixtures in order to increase the durability of concrete. the focus of this paper is on the use of special admixtures to create a durable and therefore sustainable concrete. 2. investigation and experimental set-up in an extensive research program, the influence of mix design components e.g. different admixture types, water/binder-ratio on hardened concrete properties like chloride resistance (chloride migration coefficient), sulfate resistance and water conductivity have been tested and analyzed. 2.1. testing methods the fresh concrete was assessed by the flow table spread (en 12350-5), density (en 12350-6) and air void content (en 123350-7) after mixing (5 minutes) and the flow table spread after 30 minutes. concrete cubes (150 × 150 × 150 mm) have been mad and evaluated after 28 days of curing (en 123902) according to sia 262/1 "betonbau-ergänzende festlegungen" annex a (water conductivity), b (chloride resistance) and d (sulfate resistance). limits of each test can be seen in table 1 and will be referred to in the following chapters. 558 https://doi.org/10.14311/app.2022.33.0558 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 concrete admixtures sustainable concrete test limit sulfate resistance lengthening [!] ≤ 1.0 chloride resistance chloride migration coefficient [m2/s] ≤ 10 × 10−12 water conductivity ! g/ " m2.h #$ ≤ 1.0 table 1. limits according to sia 262/1 [2, 3]. mix design a mix design b mix design c mix design d cement type cem i 42.5 cem i 42.5 cem i 42.5 cem i 42.5 cement amount [kg/m3] 300 350 300 350 w/c-ratio 0.58 0.45 0.58 0.45 max. grain size [mm] 32 32 32 32 admixture 1 dosage [%] sika® viscocrete® 0.0-0.2 sika® viscocrete® 0.6-0.8 sika® viscocrete® 0.0-0.2 sika® viscocrete® 0.6-0.8 admixture 2 dosage [%] sikacontrol® crc 1.0 / 2.0 sikacontrol® crc 1.0 / 2.0 sikacontrol® wt 0.5 / 1.0 sikacontrol® wt 0.5 / 1.0 table 2. concrete mix design. 2.2. concrete mix design the mix design has been set for all trials except the needed amount of plasticizer or superplasticizer to achieve the aimed flow table spread (fts). fts has been targeted at around 50 ± 5 cm. trials were conducted at two water/cement ratios (w/cratio): 0.58 and 0.45. the trials have been conducted with two different admixture technologies called sikacontrol® crc and sikacontrol® wt. the first one achieves an improvement of the durability of the concrete by creating a more chemical resistant concrete, whereas the second one aims on the water resistance of the concrete. sikacontrol® crc is an admixture based partly on hydroxyl combinations and one of the main effects is the in reduction in dry shrinkage and therefore shrinkage cracks. sikacontrol® wt is a water resisting admixture which incorporates mainly pore blocking and hydrophobic technologies. these concrete mix designs used in this investigation can be seen in table 2. 3. results concrete mixes with and without the special admixtures has been tested on different w/c-ratios in order to assess the performance and influence of the admixtures. limits of the tests have been listed in table 1. 3.1. influence of the admixtures at a w/c-ration of 0.58 the results of the sulfate resistance, chloride resistance and water conductivity of the concrete cubes are shown in the following figure 1 to 5. it can be observed that a significant increase of the sulfate resistance and be achieved with already 1% of the sikacontrol® crc albeit not enough to stay below the set limits. an increase in dosage up to 2% let the cubes pass the test limits (figure 1). a similar trend is shown by the results of the chloride resistance tests. a reduction of the chloride migration coefficient can be realized with 1% of the sikacontrol® crc. but again, still not good enough to pass the test. as before an increase in dosage up to 2% results in staying below the limits (figure 2). only the concrete with the higher dosage of sikacontrol® wt using the mix design c stays below the limits of the sulfate resistance test (figure 3). the chloride migration coefficient is significantly reduced versus the reference by the use of sikacontrol® wt and both dosages are below the test limits but showing only minor improvement from an increase in dosage (figure 4). sikacontrol® wt displays a significantly reduction in water conductivity versus the reference with both dosages (figure 5). 3.2. influence of the admixtures at a w/c-ration of 0.45 the results of the sulfate resistance, chloride resistance and water conductivity of the concrete cubes are shown in the following figure 6 to 9. the reference concrete with a w/c-ratio of 0.45 passes the sulfate resistance test but leaves not much room for a margin of safety. whereas both dosages of sikacontrol® crc pass the test and provide a healthy margin of safety (figure 6). all cube of the test set of mix design b pass the chloride resistance test. the positive influence of the special admixture sikacontrol® crc can be seen in the results and the reduction of the chloride migration coefficient. further investigations are necessary regarding the higher results at the higher dosage against a lower dosage (figure 7). 559 oliver schwoon acta polytechnica ctu proceedings figure 1. mix design a: sulfate resistance. figure 2. mix design a: chloride resistance. figure 3. mix design c: sulfate resistance. 560 vol. 33/2022 concrete admixtures sustainable concrete figure 4. mix design c: chloride resistance. figure 5. mix design c: water conductivity. figure 6. mix design b: sulfate resistance. 561 oliver schwoon acta polytechnica ctu proceedings figure 7. mix design b: chloride resistance. figure 8. mix design d: sulfate resistance. figure 9. mix design d: water conductivity. 562 vol. 33/2022 concrete admixtures sustainable concrete figure 10. reference concrete with various w/c-ratios: sulfate resistance. figure 11. reference concrete with various w/c-ratios: chloride resistance. the use of mix design d incorporating the admixture sikacontrol® wt will pass the test at both tested dosages and show a significant improvement against the reference concrete (figure 8). as seen before sikacontrol® wt displays a significantly reduction in water conductivity versus the reference with both dosages (figure 9). 4. conclusions the investigation has shown that the use of these special admixtures like sikacontrol® crc and/or sikacontrol® wt improve the durability of a concrete in respect to chloride and sulfate resistance and/or water conductivity. it will allow the concrete designer to choose an appropriate w/c-ratio and still maintain a margin of safety in regard to the named durability tests. furthermore, the results of the many tests conducted showed that a certain variance within the test itself has to be accounted for and a margin of safety is necessary. figure 10 displays the results of the sulfate resistance test of a reference concrete with different w/c-ratios. the results of the trialled admixtures and mix designs are shown included in figure 10 as well. it should be noted that those are only the w/c-ratios of 0.58 and 0.45 connected by a straight line. especially at higher w/c-ratios a significant positive influence can be achieved by the use of the tested admixtures. as for mix designs with lower water w/cratios a higher margin of safety can be accomplished. same conclusion as before can be drawn when it comes to the chloride resistance and the tested mix designs and special admixtures. next to well-known measure like reduction of w/cratio, use of scm etc. in order to increase the durability and therefore the service life of a structure the tested special admixtures will give concrete designers, specifiers, contractors and/or concrete producers another valuable tool to fulfil the challenging requirements of important infrastructure projects or any other needed concrete. references [1] j. schlumpf, b. bicher, o. schwoon. sika concrete handbook, zurich: sika services ag, p. 158-181, 2020. [2] the swiss standards association. sia 262/1:2013 betonbau ergänzende festlegungen, zurich: schweizer ingenieurund architektenverein, 2013. [3] the swiss standards association. sn en 206:2013+a1:2016-c1:2019, zurich: schweizer ingenieurund architektenverein, 2013. 563 https://doi.org/10.14311/app.2022.33.0463 acta polytechnica ctu proceedings 33:463–466, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague parapet structure from special shaped trc blocks jiří pazderka, petr hájek, michal nývlt∗, hana žáková czech technical university in prague, faculty of civil engineering, thákurova 7 166 29 prague, czech republic ∗ corresponding author: michal.nyvlt@fsv.cvut.cz abstract. the paper describes innovative parapet solution based on special shaped blocks made form textilereinforced concrete trc. the blocks are designed for a placement at the parapet of building with a flat roof. the solution is appropriate for buildings with precast-concrete bearing structure by following the principle of precast technology (short construction time, minimal technological pauses). however, presented parapet trc blocks can be used for energy efficient buildings as well in combination with thick layers of thermal insulation. the technical solution was registered as utility model no. cz27642(u1). keywords: concrete, parapet, prefabrication, roof, trc. 1. introduction nowadays, the emphasis is on energy savings in the field of building construction. the solution of building envelope details is important for contemporary design of buildings. an important part of this issue is the solution of the detail of flat roofs of buildings in the form of a low parapet wall. for buildings whose load-bearing structure consists of precast frame construction system, the parapet wall design is an important part of the overall design. the main advantage of precast technology is the minimization of wet processes on site which resulting to short construction time. in order to minimize the wet processes in precast construction systems, it is desirable that the parapet wall structure should also be prefabricated. however, this assumption is often not fulfilled, because at the present time the construction of the parapet wall in precast frame systems is commonly solved in the form of a masonry wall from various materials. in some cases, the parapet wall is in the form of a specially modified beam which combines the function of supporting the ceiling structure (specially adapted bearing girder) with the function of the parapet wall. however, such prefabricates usually represent a distinct architectural element (not suitable for most buildings) and at the same time almost do not consider the problem of linear thermal bridges (it cannot be used for modern energy efficient buildings). however, the requirement to minimize thermal bridges is one of the key parameters of the design of building envelope structures in today’s building construction. the solution currently used for the construction of a parapet wall on a flat roof in the form of aămasonry wall (mentioned above), presents several disadvantages. this is especially addition of another wet process into site construction which resulting to reduction of efficiency and speed of construction (not exploiting the potential of precast technology). another significant problem is the risk of deformation of the masonry parapet wall (displacement due to shear stress in masonry) over time, which is often associated with aăfailure of the roof waterproofing layer. the parapet made in the form of aăspecially modified beam (specially adapted bearing girder) is a more suitable solution in terms of in situ efficiency but these elements are always part of standardized construction systems with predefined dimensions (the possibilities of using are limited due to architectural design). another problem may be the high weight of the parapet bearing beam and the resulting complications in terms of handling the precast. 2. parapet structure from special shaped trc blocks the above-mentioned disadvantages of existing parapet structures of flat roofs of buildings based on precast frame system (masonry and parapet bearing beams) are largely eliminated using specially shaped parapet blocks made of textile reinforcement concrete (trc) [1]. the essence of the parapet block is that it is made of precast reinforced concrete with an optimized (lightweight) shape solution with aăcrosssection in the shape of the letter "l". the vertical wall of the parapet block is formed by a thin plate (30ămm thick), which is reinforced by reinforcing ribs (50×100 mm) at distances of 0.7 − 0.8 m. the stiffening ribs comprise a bending reinforcement. the low plate thickness between the reinforcing ribs is made possible using high-performance concrete (hpc) in combination with textile reinforcement (trc). the bottom horizontal slab of the block has a variable thickness which increases towards the centre of the block (thickness 30 − 80 mm). the lower horizontal plate is made of trc and, like the vertical plate, is reinforced with reinforcing ribs (50 × 100 to 137 mm) spaced 0.7 − 0.8 463 https://doi.org/10.14311/app.2022.33.0463 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en j. pazderka, p. hájek, m. nývlt, h. žáková acta polytechnica ctu proceedings figure 1. special shaped parapet block (visualization). m apart. these ribs, together with the ribs of the vertical plate, form an "l" -shaped frame structure with reinforcement in the centre of the block. the variable thickness of the horizontal slab is advantageous both in terms of precast technology (concreting will take place from the outer surface of the vertical slab) and in terms of better stability of the shaped piece (lower centre of gravity). the parapet block is equipped with a horizontal cross rib (30×100 mm) at the top, which serves for anchoring of the flashing structure (through the spacer profile). the rib on the top of the parapet block also has a reinforcing function in terms of the horizontal load on the parapet wall (wind pressure, suction, anchoring of other structures). to ensure the fixation of the position, the block is equipped with three holes in the lower horizontal plate, which are used for anchoring the block to the ceiling panel of the frame system. the shape-optimized "thin-walled" design of the parapet block is advantageous both in terms of minimizing the weight of the precast element during transportation and subsequent handling on site. at the same time, it is also advantageous the reducing of overall vertical load on the bearing structures and foundations of the building. the subtle optimized shape of the precast block using trc is also advantageous in terms of eliminating thermal bridges. only a minimal cross-section of a part of the supporting structure penetrates the thermal insulation envelope of the building. in fact, it is not a thermal bridge, because the parapet construction must always be thoroughly insulated around its perimeter. this design predetermines the parapet block for use in energy efficient buildings. for handling by lifting means, the block will be equipped with countersunk threads for screwing in the anchorages. these threads will be located on the upper surface of the horizontal reinforcing rib at the position of the vertical reinforcing ribs. 3. description of the functional principle figure 3 shows the installation of the precast parapet block (no. 1) on the ceiling structure of a precast concrete frame construction system comprising ceiling panels (no. 2) and beams (no. 3). the parapet block (no. 1) is secured (fixed) against horizontal movements by anchoring at the anchor holes (no. 4). the lower horizontal slab of the parapet block (no. 5) is"weighted" by a layer of gravity concrete (no.ă6), which thus acts as a stabilizing layer. at its lower left edge, the parapet block (no. 1) is separated from the concrete layer forming a gradient (no. 7) by means of an inserted thermal insulator strip (no. 8), which eliminates the transfer of horizontal deformations of the concrete layer due to longitudinal thermal expansion. on the upper horizontal reinforcement rib of the parapet block (no. 9) is an anchored construction of the flashing (no. 10). the vertical panel of the parapet block (no. 11) is supplemented by an additional thermal insulating layer (no. 12), which compensates for the higher thermal conductivity coefficient hpc (compared to the adjoining masonry of the external cladding (no. 13)). an additional thermal insulation layer (no. 11) is connected to the thermal insulation system of the facade cladding (no. 14), which is tightened up to the thermal insulation under the flashing (no. 15). from the inside of the roof, a vertical layer of thermal insulator (no. 16) is anchored to the parapet block. this layer follows the thermal insulation of the flat roof (no. 17). the thickness of the vertical thermal insulation (no. 15) is designed to overlap the reinforcing rib (no. 18) of the parapet block. figure 4 shows precast parapet block (no. 1) which forms the supporting structure of the parapet shown in figure 3. the parapet block (no. 1) is characterized in that it is made in the cross-sectional shape of the letter"l", where the thin vertical plate of the block (no. 11) is reinforced with ribs (no. 18). the vertical ribs are reinforced at the point of connection 464 vol. 33/2022 parapet structure from special shaped trc blocks figure 2. installation of special shaped parapet block on the building (simplified visualization). figure 3. example of parapet block installation on the ceiling structure of precast concrete frame construction system. notes: 1 parapet block, 2 reinforced concrete ceiling panel, 3 reinforced concrete beam, 4 anchoring of parapet block, 5 horizontal part of parapet block, 6 monolithic concrete layer forming a gradient and stabilizes the parapet block, 7 monolithic concrete layer forming a gradient, 8 dilatation, elastic strip, 9 the upper part of the parapet block, 10 anchoring of parapet block, 11 vertical part of parapet block, 12 lower thermal insulation layer (compensation for reinforced concrete elements), 13 filling masonry of perimeter wall, 14 main thermal insulation layer (etics), 15 insulation of parapet extruded polystyrene (xps), 16 insulation of parapet block, 17 conventional roofing layers of flat roof, 18 reinforcing rib of parapet block 465 j. pazderka, p. hájek, m. nývlt, h. žáková acta polytechnica ctu proceedings figure 4. precast parapet block. notes: 1 parapet block, 4 anchoring of parapet block, 5 horizontal part of parapet block, 9 the upper part of the parapet block, 11 vertical part of parapet block, 18 vertical reinforcement rib of parapet block, 19 horizontal reinforcement rib of parapet block, 20 rib reinforcement. to the ribs of the horizontal plate (no. 19) by triangular haunches (no. 20) and together with the reinforcing ribs of the lower horizontal plate (no. 19) form"l" frame structure. the solved parapet block (no. 1) is characterized in that it comprises a lower horizontal plate (no. 5) of variable thickness which increases towards the centre of the precast. this plate is reinforced with ribs (no. 19) which have an increasing thickness towards the centre of the prefabricated elements. the lower horizontal plate (no. 5) has three holes (no. 4) for anchoring the precast to the ceiling panel of the framework system. the parapet block (no. 1) is provided with aătransverse rib (no.ă9) at the top of the vertical plate (no. 11), which serves to anchor the parapet block flashing structure and to reinforce it in terms of horizontal loading. 4. conclusion the technical solution of the parapet block for flat roofs described in the paper was registered as a utility model no. cz27642(u1). the block is intended primarily for energy-efficient buildings based on precast frame construction systems. however, it can be used without further modifications for parapet solutions in other building construction systems (e.g. masonry buildings with precast concrete ceiling), where the advantage of the precast technology will be beneficial. the parapet block is the next element used for construction details in buildings, expanding the existing set of thin-walled precast elements, developed at the faculty of civil engineering ctu in prague [2–4]. a potential industrial partner for production is currently being searched. testing of mechanical properties is planned after the production of the first prototype series based on the results, the final dimensions of the elements for serial production will be optimized. acknowledgements this work was supported by the research project sgs19/145/ohk1/3t/11. references [1] j. pazderka, p. hájek.utility model cz27642(u1) parapet structure, especially for building platform roofs, ctu in prague, 2015. [2] j. pazderka, m. nývlt, h. žáková. underground ventilated wall based on trc blocks. iop conference series: materials science and engineering 596(1), 2019. https://doi.org/10.1088/1757-899x/596/1/012037. [3] j. pazderka, p. hájek. two innovative solutions based on fibre concrete blocks designed for building substructure. iop conference series: materials science and engineering 246, 2017. https://doi.org/10.1088/1757-899x/246/1/012046. [4] j. pazderka, e. hájková, m. jiránek. underground air duct to control rising moisture in historic buildings: improved design and its drying efficiency. acta polytechnica 57(5), 2017. https://doi.org/10.14311/ap.2017.57.0331. 466 https://doi.org/10.1088/1757-899x/596/1/012037 https://doi.org/10.1088/1757-899x/246/1/012046 https://doi.org/10.14311/ap.2017.57.0331 doi:10.14311/app.2021.30.0007 acta polytechnica ctu proceedings 30:7–11, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app plasma modification of microfibers application to lightweight cement composite containing recycled concrete jakub ďurejea, ∗, zdeněk prošeka, jan trejbala, pavel tesáreka, štěpán potockýb a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b academy of sciences of the czech republic, institute of physics, cukrovarnická 10, 162 53 prague 6, czech republic ∗ corresponding author: jakub.dureje@fsv.cvut.cz abstract. the article deals with the optimalization of composition for reinforced lightweight cement composite containing micronized recycled concrete, which will be used to produce masonry blocks. the composite material is reinforced with polypropylene microfibers. to increase the cohesion between the fibers and the cementitious matrix, the optimal modification using oxygen plasma was chosen. furthermore, a suitable foaming agent was chosen to lighten the cement matrix. a suitable ratio of cement and micronized recycled concrete was determined. finally, a cement composite was made from the optimized components. the mechanical properties of this composite were tested. the resulting mechanical properties of the lightweight samples were compared with the non-light samples. keywords: lightweight cement composite, micronized recycled concrete, oxygen plasma, plasma treatment. 1. introduction one of the most important properties of cement composites is the cohesion of the fiber with the matrix. cohesion depends on a matrix in close proximity to the fiber, on the interfacial transition zone [1, 2]. chemical and mechanical modifications of fiber surfaces are used to increase the cohesion between the fibers and the matrix. manufacturers of fibers most often apply mixtures of various chemicals to their fibers in order to increase the cohesion of the fibers with the cement matrix. however, chemical treatments are non-ecological [3]. using plasma, it is possible to modify the surface both mechanically and chemically. the process of plasma modification depends on many parameters, such as the time of modification, the working gas, the input power, etc. plasma modifies only the surface of fibers, therefore occurs to almost negligible loss of mechanical properties of the fiber itself. in addition, the process of plasma modification is more environmentally friendly than conventional methods of chemical modification [4, 5]. the most common use of recycled concrete (more than 90%) is in its unbound form as a backfill under the road. another possibility is its use in bonded form as a substitute for aggregates in cementitious composite materials. with a suitable production process and composition, cementitious composites containing recycled concrete can achieve similar mechanical properties as same composites without recycled concrete. for the production of composite material, micromilled recycled concrete was chosen, which was made from the finest fractions of recycled concrete, which is currently considered waste [6–8]. to lighten the cement composite, it is necessary to choose a foaming agent that has sufficient stability and frothiness after mixing with water and foaming. high foaming causes the formation of many air cavities in the material. the stability of the foam must ensure that the foam transfers load from the cement mixture until the material has sufficient load-bearing capacity to transfer its own weight. stability affects the resulting mechanical strength [9]. 2. materials and samples portland cement cem i 42.5 r from českomoravský cement, a.s. (radotín) was used. micronized recycled concrete was made from drainage gutters. to produce micronized recyclate was used waste recycled concrete with a fraction of 0/16 mm, which was subsequently ground by high-speed mills. the grain size of the resulting micronized recyclate was less than 0.25 mm. a suitable foaming agent was selected not only from the foaming agents intended for concrete, but also the foaming agents intended for industrial cleaning were tested. polypropylene microfibers with a diameter between 18–32 µm and a length about 12 mm were chosen to reinforce the samples. plasma modification by oxygen plasma under reduced pressure was performed using a tesla vt 214 device. microfibers were inserted into the chamber of the device 7 http://dx.doi.org/10.14311/app.2021.30.0007 http://ojs.cvut.cz/ojs/index.php/app j. ďureje, z. prošek, j. trejbal et al. acta polytechnica ctu proceedings figure 1. surface of the polypropylene microfiber, left – reference fiber, right – fiber after treatment by oxygen plasma for 480 seconds. and then air was sucked out of the chamber, after that the pressure in the chamber was only about 22 pa. subsequently, oxygen was filled into the chamber, the pressure in the chamber increased to 60 pa. subsequently, the plasma modification process was performed for 60 or 480 seconds. after completion of the plasma modification process, the chamber was filled with air to atmospheric pressure and then the fibers were removed. images of the microfiber surface before and after plasma modification process are shown in figure 1. to determine the stability of the modifications, part of the fibers were left in an environment at a temperature of about 20 ◦c and a relative humidity of 50% for 5 days (pp 480 air) before application into the cement matrix, the other fibers were applied into the cement matrix immediately after the modification. selected surface modifications of the fibers that were subsequently used to produce reinforced samples are listed in table 1. sample modification storage time [s] time [days] pp ref 0 0 pp 60 60 0 pp 480 480 0 pp 480 air 480 5 table 1. selected surface modification of polypropylene microfibers by oxygen plasma. the composition of the lightweight samples without microfibers reinforcement is in table 2. the com8 vol. 30/2021 plasma modification of microfibers set cement recyclate water w/ (c + r) foaming agent [g] [g] [ml] [−] [ml] c 1500 0 480 0.32 0 cr 750 750 480 0.32 0 cs2 1500 0 480 0.32 1.5 crs2 750 750 480 0.32 1.5 cs4 1500 0 480 0.32 3.0 crs4 750 750 480 0.32 3.0 table 2. composition of the lightweight samples without microfibers reinforcement. set cement recyclate water w/ (c + r) foaming agent fibers fibers plasma [g] [g] [ml] [−] [ml] [g] [kg.m−3] modification crf 750 750 480 0.32 0 0.75 ≈ 1 no crfp 750 750 480 0.32 0 0.75 ≈ 1 yes crs2f 750 750 480 0.32 1.5 1.43 ≈ 1 no crs2fp 750 750 480 0.32 1.5 1.43 ≈ 1 yes crs4f 750 750 480 0.32 3.0 1.875 ≈ 1 no crs4fp 750 750 480 0.32 3.0 1.875 ≈ 1 yes table 3. composition of lightweight samples reinforced by plasma treated microfibers. position of the final lightweight samples (with microfibers reinforcement) is shown in table 3. 3. experimental methods the dynamic modulus of elasticity was measured for all non-lightened samples. the resonance method was used to measure the dynamic modulus of elasticity. because it is a non-destructive method, measurement was repeat in different times (1, 7, 14, 21 and 28 days). the method is based on the measurement of fundamental resonant frequencies of transverse, longitudinal and torsional oscillations. the dynamic modulus of elasticity was measured by brüel & kjær measuring set. the measuring set contains a brüel & kjær impact hammer type 8206, a brüel & kjær acceleration sensor type 4519-003, a brüel & kjær fron-end 3560b-120 measuring control panel and a control unit. the test specimens were placed on one support in the middle when measuring longitudinal or torsional oscillations; when measuring the transverse oscillation, the samples were placed on two supports at the prescribed distances (figure 2). the response sensor (s) was glued to the prescribed position, subsequently the oscillation was excited by the impact of the impact hammer (b – exciter). the measurement was recorded by measuring control unit and then the resonant frequency for each oscillation was determined using pulse labshop software version 14.0.1. finally, the fundamental resonant frequencies were determined and dynamic modulus of elasticity and shear modulus were calculated. destructive measurement was performed on laboratory samples 40×40×160 mm. each set of samples figure 2. diagram of the sensor location (s) and hammer impact point (b) when measuring longitudinal (left), transverse (middle), and torsional (right) oscillations using the resonance method [10]. contained 6 test specimens. samples were stored for 28 days into water basin in a laboratory environment at 22 ± 1 ◦c. first, three-point bending tests were performed using a hydraulic press, during which the test specimens were halved. then compressive tests were performed on the halves of the test specimens (approximately 40 × 40 × 80 mm) using a hydraulic press. 4. results and discussion to select the optimal composition of the cement composite material, the individual components were first selected during testing on non-light samples, then the components were applied to a lightweight matrix. the polypropylene fibers were modified by oxygen plasma. the plasma modification of the fiber surface has a positive effect on the mechanical properties of the tested cementitious composite containing micronized recyclate. during plasma modification process the fibers lose almost no weight, which proves 9 j. ďureje, z. prošek, j. trejbal et al. acta polytechnica ctu proceedings figure 3. flexural strength of samples containing fibers with different plasma modification times. figure 4. compressive strength and density of samples with different amount of foaming agent. figure 5. compressive strength and density of lightweight cementitious composites containing recycled concrete reinforced by plasma modified fibers. 10 vol. 30/2021 plasma modification of microfibers that the diameter of the fibers is almost not reduced and mechanical properties of fibers are approximately the same as before the modification. when examining the surface of the fibers by sem, the mechanical effect of the plasma modification on the surface was noticeable only for the fibers that were modified for 240 seconds or more. the mechanical properties of the test specimens were best for samples containing fibers modified with oxygen plasma for 60 seconds, with longer modification times the mechanical properties were no longer improving (figure 3). oxygen plasma modification for 60 seconds was chosen as the optimal modification for polypropylene microfibers. the optimal amount of recycled concrete was chosen based on previous experiments [11, 12]. the ratio between cement and micronized concrete recycled was determined to 1:1. a total of six different foaming agents were tested to lighten the samples. the best results for lightening the cement matrix containing micronized recycled concrete were achieved by the foaming agent sika lightcrete 400. the most lightened samples had a bulk density about 800 kg.m−3 and a compressive strength about 4 mpa (figure 4). finally, lightweight samples reinforced with plasma-modified fibrous reinforcement were tested. the amount of fibers was chosen based on the recommended dosage from the manufacturer, the samples contained one kg of fibers per m3 of cement composite. at the recommended amount of fibers, samples containing plasma-modified microfibers achieved almost the same flexural strengths as samples containing fibers without plasma modification. samples containing plasma modified fibers achieved slightly higher compressive strengths compared to samples without plasma modification. however, the increase in compressive strength was not very significant. the results are shown in figure 5. 5. conclusion plasma modification by oxygen plasma increases the cohesion of polypropylene microfibers with the cement matrix. the mechanical effect of plasma modification by oxygen plasma on the fiber surface is insufficient. when dosing micronized concrete recycled and cement into the matrix in a ratio of 1:1, good mechanical properties of the cement composite material are achieved. when the microfiber reinforcement was applied to the lightweight cementitious matrix, there was an increase in bulk density compared to the same lightweight material without fibrous reinforcement. the use of oxygen plasma to modify fibers surface and their application into a lightweight cement matrix does not have a significant effect on the mechanical properties of this lightweight material. in order to achieve better mechanical properties of this lightweight material, it is necessary to increase the amount of fibers and to choose another type of plasma modification. to achieve better mechanical properties of the reinforced lightweight cement matrix, it may be appropriate to modify fibers surface by argon plasma. the modification by argon plasma is expected to have a significantly higher mechanical effect on polypropylene fibers surface compared to the oxygen plasma modification. 6. acknowledgements this work was financially supported by the czech technical university in prague – the project sgs19/148/ohk1/3t/11. the authors also thank to the center for nanotechnology in civil engineering at the faculty of civil engineering, czech technical university in prague. references [1] a. bentur, s. mindess. fibre reinforced cementitious composites. new york, 2007. [2] j. p. ollivier, j. c. maso, b. bourdette. transition zone in concrete. advanced cement based materials 2(1):30–38, 1995. doi:10.1016/1065-7355(95)90037-3. [3] m. m. kabir, h. wang, k. t. lau, f. cardona. effects of chemical treatments on hemp fibre structure. applied surface science 276:13–23, 2013. doi:10.1016/j.apsusc.2013.02.086. [4] s. r. sabreen. cold gas plasma surface modification ű optimize plastics bonding adhesion, 2012. [5] j. tinneke, m. rino, n. geyter, et al. plasma surface modification of biomedical polymers: influence on cell-material interaction. plasma chemistry and plasma processing 32:1039–1079, 2012. doi:10.1007/s11090-012-9394-8. [6] v. w. y. tam, c. m. tam. a review on the viable technology for construction waste recycling. resources, conservation and recycling 47(3):209–221, 2006. doi:10.1016/j.resconrec.2005.12.002. [7] i. vegas, j. a. ibanez, a. lisabona, m. cortazar, a. s. abd frias. pre-normative research on the use of mixed recycled aggregates in unbound road sections. construction and building materials 25(5):2674–2682, 2011. doi:10.1016/j.conbuildmat.2010.12.018. [8] z. prošek, p. tesárek, j. trejbal. recycling and its use in concrete waste processing by high-speed milling. acta polytechnica ctu proceedings 21:28–32, 2019. doi:10.14311/app.2019.21.0028. [9] k. ramamurthy, e. k. nambiar, g. i. s. ranjani. a classification of studies on properties of foam concrete. cement and concrete composites 31(6):388–396, 2009. doi:10.1016/j.cemconcomp.2009.04.006. [10] čsn 73 1372 nedestruktivní zkoušení betonu–rezonačnní metoda zkoušení betonu, 2012. [11] z. prošek, j. trejbal, v. nežerka, et al. recovery of residual anhydrous clinker in finely ground recycled concrete. resources, conservation and recycling 155, 2020. [12] z. prošek, j. topič, j. ďureje, j. trejbal. srovná́ní vlivu mikromletého betonu a mramorové moučky na mechanické vlastnosti cementových past. in waste forum 2:262–267, 2018. 11 https://doi.org/10.1016/1065-7355(95)90037-3 https://doi.org/10.1016/j.apsusc.2013.02.086 https://doi.org/10.1007/s11090-012-9394-8 https://doi.org/10.1016/j.resconrec.2005.12.002 https://doi.org/10.1016/j.conbuildmat.2010.12.018 https://doi.org/10.14311/app.2019.21.0028 https://doi.org/10.1016/j.cemconcomp.2009.04.006 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.35.0014 acta polytechnica ctu proceedings 35:14–18, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague modern railways for the 21st century – a challenge not only for designers josef kolář czech technical university in prague, faculty of mechanical engineering, department of automotive, combustion engines and railway engineering, technická 4, 160 00 prague 6, czech republic correspondence: josef.kolar@fs.cvut.cz abstract. the paper outlines the basic goals for modern rail transport in the czech republic. this contribution points out some challenges and technical tasks that project engineers and operators of the rail transport in the czech republic will have to deal with if they want to succeed in the conditions of the liberalized european railway market after the year 2035. moreover, the article assesses the demands placed on safety of the railway transport and presents suggestions for reduction of the consequences of a rolling stock-car collisions in the regional transport. keywords: active safety, design vehicles, etcs, high speed railway, modern railway transport, passive safety, regional railways. 1. introduction in the 21th century, modern railways must not be just an interconnected and interoperable network of european high-speed and corridor lines, e.g. see figure 1. it must represent a highly ecological and safe transport system, which offers customers fast and reliable transport throughout europe by its services and quality. the transport system, which will be more timeand cost-effective and acceptable than competitive air or individual car transport (ict) for people and goods transportation to a distance of about 1000 km. modern rail transport in the czech republic has to meet the following basic goals: • to be a leader in the safe, attractive and barrier-free passenger transport, not only in the long-distance transport, but also in the regional and suburban transport, • with the quality services offered to passengers it has to offer a higher level of connectivity to the regional passenger transport to junctions (stations) and thus to contribute to a reduction of the part of the individual car transport in passenger transport for work or culture. the model of ecological, time-bound and affordable mobility needs to be strengthened, • it has to offer operationally safe, energy-efficient and low-emission (noise, co2, nox , dust particles) transport of goods and freight, which will be more competitive and more environmentally friendly than individual truck transport, • it must significantly increase its share on the international container transport and the european combined transport (road – railway), because this system is able to significantly decrease an accident rate of truck transport. trucks should be mainly used for the first and the last mile freight transport from the railway junction (modern transshipment point) to the customer and back, not for transport across the whole country or the central europe, • within the freight transport it has to ensure the ecological collection of municipal waste to selected large-capacity incinerators and sorting plants, because by this way it is possible to meet the objective of the ban on landfilling of municipal waste and its secondary use. by taking these steps with the high level of "traffic electrification" the modern railways, while maintaining a certain level of mobility, can make a significant contribution to meet the emission reduction targets defined in the eu program documents called as "white book of transport". 2. challenges not only for rolling stocks project engineers to meet these goals requires: • significant investments costs in the infrastructure (building hsr, modernization of the contemporary railways and a completion of electrification – transition to 25kv/50hz ac traction system, modernization of railway network junctions – transfer points from the regional railway to hsr, loading hubs for european combined transport. • extensive investments in vehicles (purchase of new rolling stocks, new high-speed electric multiple units for speed of 300 kph, cost-effective fleet modernization, etcs implementation, dual-powered vehicles for regional transport, new vehicles for combined transport, . . . ). • change of transport models and the passenger transport in order to achieve maximum and efficient 14 https://doi.org/10.14311/app.2022.35.0014 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 35/2022 modern railways for the 21st century figure 1. a trans-european rail lines ten-t [1]. utilization of new high-speed lines and vehicles. • new models of mobility in the area of the central europe and a thinking and behaviour change of the people and passengers in favour of safe and environmental-friendly transport. only the rapid fulfilment of these points and elaborated training of new railway experts and employees is able to prepare czech railway for a great competition in the conditions of the liberalized european railway market after 2035. achieving of low operating costs while meeting high level of operational safety and environmental requirements will be key factors, which will decide about the success or failure of our operators on this market. figure 2. high speed agv or ice 4 for passenger transport in czech republic? [1]. 2.1. high-speed railway transport versus corridor lines = new transport models the requirements for a high safety level of high-speed railway traffic on hsr are embodied in the technical specifications for interoperability (tsis) for the design of railway tracks and rolling stocks. tsi hs fulfilment and european train control system etcs implementation (level 2 and level 3) shows that the hsr system, with the appropriate operational load of high-speed units, significantly contributes to economic and environmental mobility and demonstrate competitiveness against air or road transport. the czech republic has to significantly accelerate building of hsr if it wants to offer an economic and fast connection between northern and southern europe to european operators after the year 2035. it would be inappropriate if the czech republic is bypassed and the international transit passes through the federal republic of germany and austria. the transfer of transport capacity in the international and the domestic long-distance passenger transport on high-speed railways with maximum axle load of 17 tonnes and speed limit up to 350 kph will require a change in transport models of regional services and transfer nodes. daily operation between approx. 5-23 hours can be realized by electric multiple units with distributed power, see figure 2. for night bed and lounger express trains with operation from about 9 pm to 7 am, it is more appropriate to use units with concentrated traction power in front and rear car with regard to passenger comfort. for hsr operation these multi-system drive head cars should not exceed the mass of 67 tonnes. at this point it is necessary to point to the fact that czech operators will have to buy these vehicles in near future. 15 josef kolář acta polytechnica ctu proceedings these new transport models should free up existing railway corridors for a high-speed regional transport and for international container and combined freight transport. for conventional freight-trains sž s. p. (railways administration s.e.) has to ensure suitable and efficient use of incline and directionally more demanding railway line kolín – kutná hora – havlíčkův brod – křižanov – brno, see figure 3. however, modern railways have to maintain a high level of operational safety when meeting the new requirements. figure 3. railway transit corridors in czech republic. 2.2. modern regional railway, safety versus railway ecology the safety of rolling stocks operation is divided into two groups: • active safety – represents elements and systems to prevent accidents of rolling stock among themselves or with road users. • passive safety – represents the components and systems designed to reduce the consequences of an accident. 2.2.1. modern speed regional railway for railway lines with a speed limit above 160 kph, level crossing with road infrastructure is required. this requirement of separation of the railway line from the road provides a sufficient level of active safety. in the field of passive safety of rolling stock, new requirements for the strength (en 12 663) and crash resistance of rolling stock (en 15 227) were defined at the end of the 20th century, see train avg – figure 4. the implementation of these new measures, together with the growing demands on multi-system (three-system, dual-powered or hybrid) vehicles, accessibility and better equipment of vehicle interiors (wifi, better information systems, noise reduction, etc.) is reflected in an increase of vehicle body weights. figure 4. visualization of energy absorption device – hsr train unit agv [2]. the manufacturers are trying to compensate for this increase in weight with a light carbody design and with a change in the concept of vehicle running gear. instead of bogies with outer frames, concepts with an inner frame are being developed. but even these solutions have their design limitations (eg. higher sensitivity to changes of wheel load forces, less space for installation of the wheelset drive, reduced angular stiffness of the primary and secondary suspension, which leads to utilization of torsional stabilizers, and others). 2.2.2. regional and suburban railway transport versus road transport the situation is somewhat more demanding in terms of safety and ecology of operation in the regional transport. the requirement for a regional rolling stock to meet the requirements of en 12 663 – category p2, en 15 227 – category ci represents an increase in its weight and the resulting higher rolling resistance, greater inertia effects and thus higher energy consumption. modern 21st-century regional railways must therefore use rolling stocks with electric power transmission, as only this type of drive makes possible to achieve a high percentage of vehicle kinetic 16 vol. 35/2022 modern railways for the 21st century figure 5. bogie siemens sf 7500 with inner frame and inside bearings – train unit mireo [3]. energy recovery with the consistent application of electrodynamic brakes on all vehicle wheelsets. figure 6. electric regional unit mireo [3]. regional railway lines quite often cross roads in the form of railway crossings. there are about 8,000 railway crossings in the czech republic with varying degrees of security, in [4]. any unprotected level crossing increases the risk of crash between regional train and road vehicle. compliance with the requirements of standards en 12 663 (category p2), en 15 227 (category ci) makes the railway vehicle a clear winner in the case of a collision with car. nowadays, most collisions of rolling stocks with the car leads to a significant car body destruction and serious injuries or death to car crew members. in this context, the question arises for a broader expert discussion: should we equip vehicles for regional transport with additional increased weight (resulting from the installation of the additional passive safety elements and the requirement for controlled deformation of vehicle fronts) and thus increase vehicle weight, extend starting and braking distances, increase consumption and reduce competitiveness with bus transport? meeting the requirements of standards en 12 663 (category p2), en 15 227 (category ci) de facto means that economically operated and ecologically advantageous, light two-axle motor vehicles, called as rail buses, will disappear from regional railways. would it not be more appropriate to require regional rolling stock operated on local railways to meet the requirements of en 12 663 category p3, en 15 227 – category cii and to equip both types of vehicles (rail and road) with active safety features that would significantly eliminate road vehicle collisions with a light rail vehicle (rail bus)? figure 7. design study of the rail – electric bus [5]. the potential to reduce the consequences of a collision between a rail vehicle and cars is possible: > in the field of active safety, it is appropriate to consider and use the possibility of communication of the mobile part of etcs placed in new rolling stocks with the navigation and control system of cars. when all rolling stocks are equipped with an etcs safety guard, which allows the vehicle ‘s position to be monitored on the track, it is possible to link the rolling stock communication with the level crossing or directly with the motor vehicle. thanks to the exchange of information about the railway vehicle drive, car control units could control the driver’s activity and possibly directly prevent the driver from driving over the railway crossing and thus prevent an accident. > in the field of passive safety, it is appropriate to assess the possibility and reality of the application of new safety features ("protective frames or panels") on the fronts of regional rolling stocks. these more flexible protective frames, which the driver would "shoot" in 17 josef kolář acta polytechnica ctu proceedings front of the rail vehicle in the event of an immediate avoidance of the accident. these forward-moved safety elements will absorb and "soften" the impact of the front of the light rail vehicle to the side of the car and reduce the penetration of the coupling into the passenger zone. figure 8. revolution very light rail [6]. 3. conclusions the requirement to implement etcs-level 2 will certainly make a significant contribution to a higher level of active safety in the regional rail traffic. however, the next, entirely logical step should be to start intensive cooperation between etcs creators and car navigation and control system manufacturers. the transfer of information about train’s position to the car navigation and its display to the driver and storage in a "black box" recording device can significantly increase the responsibility of car drivers for their behaviour at unprotected level crossings. this cooperation of the etcs mobile part with road vehicles will be necessary for autonomous vehicles. figure 9. very light rail for export to turke [7]. acknowledgements this research has been realized using the support of the technology agency of the czech republic, programme national competence centres, project #tn01000026 josef bozek national center of competence for surface transport vehicles this support is gratefully acknowledged. references [1] j. kolář. vysokorychlostní tratě a doprava v čr. [unpublished lecture], fakulta dopravní čvut v praze, 2020-11-10. [2] https://www.alstom.com,https://fdocuments.in/ document/alstom-english-agv.html. [3] https://www.mobility.siemens.com/global/en/ portfolio/rail/rolling-stock/ commuter-and-regional-trains.html. [4] j. kolář, j. seidl, v. zelený, o. vincik. statistika nehodovosti tramvají a vozidel regionální železnice. [research report], praha: ctu fme, department of automotive, combustion engine and railway engineering, 2020. report no. z20-03 u 12 120 fs čvut. [5] j. kolář. konstrukční studie částečně nízkopodlažního kolejového elektrobusu. in zborník prednášok i. – xxiv. medzinárodná konferencia – súčasné problémy v koľajových vozidlách. žilina: vts pri žu, 2019. p. 317–324, isbn 978-80-89276-58-5. [6] https://warwick.ac.uk/fac/sci/wmg/research/ hvmcatapult/research/rail/vlr. [7] https://www.transportxtra.com/publications/ local-transporttoday/news/47689/ uk-manufacturer-unveils-very-light-rail-vehicle, https://www.severn-lamb.com/images/products/ ulr-express/dsc_4741.jpg. 18 https://www.alstom.com, https://fdocuments.in/document/alstom-english-agv.html https://www.alstom.com, https://fdocuments.in/document/alstom-english-agv.html https://www.mobility.siemens.com/global/en/portfolio/rail/rolling-stock/commuter-and-regional-trains.html https://www.mobility.siemens.com/global/en/portfolio/rail/rolling-stock/commuter-and-regional-trains.html https://www.mobility.siemens.com/global/en/portfolio/rail/rolling-stock/commuter-and-regional-trains.html https://warwick.ac.uk/fac/sci/wmg/research/hvmcatapult/research/rail/vlr https://warwick.ac.uk/fac/sci/wmg/research/hvmcatapult/research/rail/vlr https://www.transportxtra.com/publications/local-transport today/news/47689/uk-manufacturer-unveils-very-light-rail-vehicle https://www.transportxtra.com/publications/local-transport today/news/47689/uk-manufacturer-unveils-very-light-rail-vehicle https://www.transportxtra.com/publications/local-transport today/news/47689/uk-manufacturer-unveils-very-light-rail-vehicle https://www.severn-lamb.com/images/products/ulr-express/dsc_4741.jpg https://www.severn-lamb.com/images/products/ulr-express/dsc_4741.jpg acta polytechnica ctu proceedings 35:14–18, 2022 1 introduction 2 challenges not only for rolling stocks project engineers 2.1 high-speed railway transport versus corridor lines = new transport models 2.2 modern regional railway, safety versus railway ecology 2.2.1 modern speed regional railway 2.2.2 regional and suburban railway transport versus road transport 3 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0058 acta polytechnica ctu proceedings 33:58–62, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague durability of gfrp connectors under sustained compression load for use in sandwich walls stefan carstens∗, matthias pahn technische universität kaiserslautern, department of civil engineering, erwin-schrödinger-straße 52, 67663 kaiserslautern, germany ∗ corresponding author: stefan.carstens@bauing.uni-kl.de abstract. precast concrete sandwich panels are used to fulfil the rising thermal requirements. the sandwich walls consist of three layers, a facing, a thermal insulation layer, and a load bearing layer. the two outer layers are coupled by connectors made of glass-fibre reinforced polymer. a lack of knowledge about load-bearing behaviour prevents the removal of sustained compressive loads. in the context of this article, tests under sustained compressive load are presented. to represent closely the in service-conditions of sandwich walls, the examined connectors were subjected to a saturated alkaline concrete environment as well as to a specified stress level till failure occurs. thus, the experimental setup combines alkaline resistance and creep rupture tests into one comprehensive testing. by using temperature effects as an accelerating factor, reasonable test durations were enabled. the obtained time to failure line was determined to extrapolate the characteristic values of the long-term strength for a service life up to 50 years. the test results are compared and evaluated with existing test results under a sustained tensile load. keywords: durability, gfrp connector, sandwich wall sustained compression load. 1. introduction precast concrete sandwich panels have been used as exterior wall systems for many years and have proven themselves in practice [1]. they are characterised by numerous design options (fig. 1) as well as high resistance to weathering and ageing [2]. the prefabricated wall elements have a three-layer structure consisting of a facing layer, a factory-installed thermal insulation layer, and a load-bearing layer. in standard cases, the facing and load-bearing layers are made of normal concrete [3] and in individual cases of lightweight concrete [4] or high-performance concrete [5]. the facing layer serves as an architectural design element that is at least 70 mm thick [6] and fulfils the function of weather protection for the thermal insulation material. the load-bearing layer performs load-bearing functions and is at least 140 mm thick. the thermal insulation materials used are expanded polystyrene (eps) according to din en 13163 [7], extruded polystyrene (xps) according to din en 13164 [8], and insulation materials based on polyurethane (pur) according to din en 13165 [9]. usual insulation thicknesses are between 60 to 240 mm. discrete connectors made of corrosion-resistant materials are used to connect the facing and the load bearing layer [6]. these must allow deformations of the facing layer due to thermal effects as free as possible from constraining forces and simultaneously transfer wind and dead-loads across the thermal insulation [10]. to avoid thermal bridges, metallic connectors such as stainless-steel connectors are increasingly being replaced by connectors made figure 1. precast concrete sandwich panels for use as a façade [5]. out of fibre-reinforced plastic (frp). the connectors usually consist of a thermosetting plastic matrix and unidirectional reinforcing glass fibres. the individual components perform different tasks within the composite [11]. the fibres used are thin (! 3 . . . 25 mm), unidirectional oriented continuous glass fibres, which are completely enclosed by a synthetic resin matrix. due to their mechanical properties, the inorganic fibres have a decisive influence on the stiffness and strength of the composite material [12]. thermosetting polymers such as epoxy resins (ep), vinyl ester resins (ve), and vinyl ester urethane resins (veu) are mainly used as materials for the synthetic resin matrix. the matrix fixes the fibres in the desired geometric position, transfers the stress between the fibres, prevents the fibres from buckling under a compressive load, and protects the fibres from external mechanical and chemical stress. the advantages of the individual components are combined by the interaction of the glass fibre and the polymer matrix. in addition to the properties of the components and their proportions, pultrusion 58 https://doi.org/10.14311/app.2022.33.0058 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 durability of gfrp connectors name isolink thermopin tm-mc anchor matrix material veu ep ve fibre material wt. % 87.1 n/a 74.0 vol. % 74.3 n/a 53.0 surface ripped wound flat min. tensile strength n/mm2 1.000 1.500 700 tensile young’s modulus n/mm2 60.000 60.000 40.000 section dimensions mm !c 12.0; !e 13.5 !e 7.3 h/w 9.8/5.7 table 1. characteristics of gfrp connectors acc. to [13]. [12] and the specific interaction of matrix resin and reinforcing fibre are decisive for the properties of the composite material. connectors made of glass-fibre reinforced plastic differ significantly in terms of material composition and cross-sectional geometry. table 1 gives an overview of the material and geometry characteristics of different grp connectors. 30° ø �1 2� ø �1 3. 5� figure 2. the geometry of the examined connector’s head made out of glass fibre reinforced polymers. in the further course of this paper, the isolink [14] shown in figure 2 will be considered as the connector for sandwich wall elements. the connector has a core diameter of 12 mm made out of glass fibre reinforced polymers. the total diameter of the connector, core, and ribs, is 13.5 mm. the ribs are made by indented trapezoidal thread with a profile depth of 0.6 to 0.75 mm and a pitch of 8 mm. the ends of the connector are bevelled at an angle of 30◦. in addition to temporary loads due to wind and temperature effects, the connectors must also be withstanding permanent loads due to the dead load of the facing layer. loads acting vertically on the facing layer are carried by horizontally arranged anchors. the dead load of the facing layer is transferred using load-bearing anchor systems consisting of horizontal and diagonal anchors. the support anchor system consisting of horizontal and diagonal anchors and the associated static system for calculating the distribution of forces in the support anchor system under the dead load of the facing layer are shown in figure 3. as a result of the force-deflection, tensile forces arise in the diagonal connector nt,sust and compressive forces in the horizontal connector nc,sust which have a permanent effect. 2. load-bearing behaviour under short-term compression load while the load-bearing behaviour of the connectors under short-term [15] and long-term tension load [16] is known, there is a lack of knowledge about the loadbearing behaviour under compressive load, especially the stability load-bearing behaviour under consideration of exposures (humidity, temperature and alkalinity). in the context of this paper, therefore, the results of experimental investigations into the loadbearing behaviour under compressive load with large insulation thicknesses are presented. during the test, the connectors are subjected to short-term and longterm pressure loads. in evaluate the load-bearing behaviour under short-term compressive loading, compression tests were carried out at insulation thickness of 160 mm, 250 mm, 300 mm and 350 mm [17]. the specimen geometry is shown in figure 4. to ensure a realistic anchorage of the connectors in the facing and load bearing layer, the connectors are embedded in normal concrete. the thickness of each concrete layer is 120 mm, the other cross-sectional dimensions are 300 × 300 mm. the core layer consists only of the connector, no thermal insulation is used. the test specimen is loaded during the test procedure using a servo-hydraulic test cylinder controlled by displacement. the transverse speed is 0.5 mm/min. simultaneously, the force is measured using a load cell and the deformation through inductive displacement transducers. the determined loads are shown in figure 5. with an insulation thickness of 160 mm, the connectors fail locally in the clamping area combined on thrust and pressure. with other insulation thicknesses, an interlaminar shear crack occurs over the entire length of the connector. furthermore, figure 5 shows the recalculation of the stability tests according to equation 1. a doublesided clamping of the connector is assumed (euler case iv, k = 0.5). the second moment of inertia is using the core diameter !c = 12 mm without considering a stabilizing effect of the ribs. pcr = π2ei (kl)2 (1) 59 stefan carstens, matthias pahn acta polytechnica ctu proceedings figure 3. support anchor system (left) and static system (right) to calculate the distribution of forces in the support anchor system due to the dead load of the facing layer ved. figure 4. geometry of the test specimens for shortand long-term tests under compression load. it can be shown that the results of the experimental investigations for the insulation thicknesses 250 mm, 300 mm and 350 mm correspond very well with the recalculation. with an insulation thickness of 160 mm, the load-bearing capacity is overestimated by the recalculation. the reason for this is the change in the failure mode described above. 3. experimental investigation under sustained compression load the mechanical and physical properties of the composite material can degrade over time due to physical effects from e.g. creep and environmental influences. since the transfer of the dead load of the facing layer must be ensured over the entire service life of the façade, these influences must be taken into account when determining permissible load-bearing capacities. the connectors are particularly exposed to temperature and humidity. furthermore, the alkalinity of the concrete can damage the matrix. figure 5. experimentally and numerically determined the load-bearing capacity of the connector under compression load as a function of the insulation thickness. to evaluate the load capacity of the connectors under compressive load, creep tests to failure are carried out to determine the respective service life. 3.1. experimental set-up and execution of experiments under sustained load the test setup for the creep tests is shown in figure 6. the test specimens described above are placed in a tempered water bath. the water level reaches to just below the upper edge of the lower concrete abutment. the water is held at 60 řc to simulate increased surfaced temperatures in the area of the façade. the increased temperature causes water vapour to form, which settles on the connector and loads to a moderate moisture load. the force is applied by a centrally arranged hydraulic cylinder. the hydraulic pressure in the cylin60 vol. 33/2022 durability of gfrp connectors 1 23 4 figure 6. experimental set-up for compression tests under continuous load and exposure to humidity, temperature and alkalinity; 1 hydraulic jack and oil pressure sensor; 2 test specimen; 3 temperature regulation; 4 stainless-steel profiles for guiding the specimen. der is applied by a manual hydraulic pump and measured by an oil pressure sensor. the deformations of the specimen are measured simultaneously by an inductive displacement transducer. the concrete abutments of the test specimens are guided in parallel by a stainless-steel construction to prevent later deflections. 3.2. results of experimental investigations under sustained load the time-deformation diagram in figure 7 shows the course of the deformation over the load duration. the three technical creep ranges are marked. the elastic deformation of the test specimen can be seen at the beginning. in the initial area, the test specimen shows an increase in linear deformation (primary creep area). the test specimen is then subjected to constant creep (secondary creep). this is followed by an increasing creep deformation (tertiary creep) which is finally followed by the failure of the test specimen. compared to the load capacity of 18.9 kn shown in figure 5, a reduction of the load capacity to 16 kn can be determined for a service life of approximately 100 h. this reduction is probably due in particular to the influence of temperature and humidity. 3.3. evaluation of the results of experimental investigations further service lives are achieved by varying the loads. figure 8 shows the previous service lives of running and finished tests and the associated loads. with decreasing load, an increase in service life is expected. from the service lives of the individual tests figure 7. deformation of the test specimen over the test period at an insulation thickness of 350 mm. and loads, a permissible load for the intended service life of 50 years is then to be derived by extrapolation. the extrapolation should be performed according to [18]. the calculated load capacity for an insulation thickness is on the safe side as insulation thicknesses range from 60 to 350 mm. figure 8. load time to failure diagram of creep tests at an insulation thickness of 350 mm. 4. conclusions multilayer concrete sandwich panels have been used as exterior wall systems for many years and have been proven themselves in practice. connectors made of glass fibre reinforced polymers are increasingly being used to couple the concrete layers. to transfer the dead load of the facing layer through a support anchor system consisting of connectors, the connectors must permanently absorb compressive loads. in this paper, experimental investigations under short-term and long-term compression load are presented. it can be shown that the tests with insulation 61 stefan carstens, matthias pahn acta polytechnica ctu proceedings thicknesses between 250 and 350 mm can be recalculated very well using the euler equation. in the longterm tests, an increased influence of temperature and humidity on the failure load can be derived from the initial results. finally, a concept for extrapolation is proposed for the determination of permissible load capacities. acknowledgements the authors thank schöck bauteile gmbh for the financial and technical support provided. references [1] m. pahn. beitrag zur ermittlung von schnittund verformungsgrößen bei mehrschichtigen stahlbetonwandtafeln mit verbindungsmitteln aus glasfaserverstärktem kunststoff kaiserslautern germany, ph.d. thesis, kaiserslautern, deutschland, 2011. [2] r. gastmeyer. neuentwicklungen bei der konstruktion mehrschichtiger stahlbeton-wandtafeln mit integrierter wärmedämmung. betonund stahlbetonbau 98(6):3608, 2003. https://doi.org/10.1002/best.200301750. [3] m. pahn, j. schnell. einfluss der verbundtragwirkung bei mehrschichtigen stahlbetonwandtafeln mit innen liegender wärmedämmung. betonund stahlbetonbau 106(8):551-60, 2011. https://doi.org/10.1002/best.201100040. [4] f. hanz. beitrag zur bemessung von dreischichtigen wandtafeln mit deckschichten aus haufwerksporigem leichtbeton kaiserslautern, ph.d. thesis, kaiserslautern (deutschland), 2017. [5] f. müller, c. kohlmeyer, j. schnell. sandwichelemente mit deckschichten aus hochleistungsbeton und einem kern aus extrudiertem polystyrol. betonund stahlbetonbau 110(12):799-810, 2015. https://doi.org/10.1002/best.201500038. [6] din en 1992-1-1/na 2013. nationaler anhang national festgelegte parameter eurocode 2: bemessung und konstruktion von stahlbetonund spannbetontragwerken teil 1-1: allgemeine bemessungsregeln und regeln für den hochbau, 2013. [7] en 13163 2016. thermal insulation products for buildings factory made expanded polystyrene (eps) products specification, 2016. [8] en 13164 2015. thermal insulation products for buildings factory made extruded polystyrene (xps) products specification, 2015. [9] en 13165 2016. thermal insulation products for buildings factory made rigid polyurethane foam (pu) products specification, 2016. [10] p. bindseil. stahlbeton fertigteile nach eurocode 2 konstruktion, berechnung, ausführung, werner verlag, 2012. [11] a. schumann, m. may, m. curbach. carbonstäbe im bauwesen. betonund stahlbetonbau 113(12):868-76, 2018. https://doi.org/10.1002/best.201800077. [12] h. schürmann. konstruieren mit faser-kunststoff-verbunden, springer verlag, 2017. [13] dafstb 615 (en). commentary to en 1992-4 design of fastenings for use in concrete, 2019. [14] zulassung z-21.8-1894. schöck isolink für mehrschichtige betontafeln, 2018. https://www.dibt.de/de/service/ zulassungsdownload/detail/z-218-1894. [15] f. micelli, a. nanni. mechanics of composite materials 39(4):293-304, 2003. https://doi.org/10.1023/a:1025638310194. [16] m. l. keller, m. pahn. durability of gfrp bars with different bar diameters. iabse symposium, guimaraes 2019: towards a resilient built environment risk and asset management, p.603-10, 2019. https://doi.org/10.2749/guimaraes.2019.0603. [17] bericht 06046pa/528. untersuchungen zur eignung von combar hohlwandankern aus glasfaserverstärktem kunststoff als verbindungselement von kerngedämmten wänden mit aufstehender vorsatzschale unreleased bericht 10012pav/14511 2015 versuche zur bestimmung des widerstandes gegen stabilitätsversagen von schöck thermoankern (unreleased), 2008. [18] a. weber. prüfkonzepte für bewehrungsmaterialien mit zeitabhängigen widerständen bauingenieur, 2018. 62 https://doi.org/10.1002/best.200301750 https://doi.org/10.1002/best.201100040 https://doi.org/10.1002/best.201500038 https://doi.org/10.1002/best.201800077 https://www.dibt.de/de/service/zulassungsdownload/detail/z-218-1894 https://doi.org/10.1023/a:1025638310194 https://doi.org/10.2749/guimaraes.2019.0603 doi:10.14311/app.2021.30.0131 acta polytechnica ctu proceedings 30:131–134, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app modelling of time-dependent behaviour of particulate thermoset polymers jan vozáb∗, jan vorel czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jan.vozab@fsv.cvut.cz abstract. a preliminary study of a numerical model describing the behaviour of polymer-based composites is presented. the numerical model consists of three main parts. the first is the microplane m4 model, which is the main part of the model and is used to simulate elastoplastic behaviour and damage. the second part consists of a generalized maxwell model, which adds the effect of linear creep of the material to the calculation. the last part is a free volume model that extends the linear creep to the nonlinear creep. the creep is calculated on the deviatoric part of the normal stress of each microplane, which allows the model to capture the polymer behaviour adequately without adjusting the free volume of the model. keywords: creep, free volume, microplane, thermosets, viscoelasticity. 1. introduction many different materials are used in building constructions. one specific type that is increasingly used is epoxies. these polymers are often used to form different types of composites, as universal adhesives, rigid foams, or as structural adhesives [1]. like many polymeric materials, they are subject to nonlinear creep, which is associated with susceptibility to changes in temperature and humidity [2–4]. to properly design components made of this material, the appropriate numerical tool is essential. many models, describing broad nonlinear time-dependent responses of these material types, have been developed [5, 6]. by combining these approaches with the generally accepted material model for mechanical loading, the overall material response can be adequately described. in this paper, we introduce the novel approach combining a microplane material model m4 [7, 8], which is enhanced with the free volume approach on the microplane level to capture the time-dependent behaviour [5]. 2. numerical model simulating nonlinear behaviour is challenging because no nonlinear theory has been introduced that can describe all materials. however, many models have been developed that describe broad nonlinear responses [5, 6]. by combining these approaches, it is then possible to describe the behaviour of these materials. the numerical model described in this paper consists of three main parts. the microplane material model m4 [7, 8] with improvements presented in [9] is used as a fundamental part. the maxwell chain on the microplane level is used to capture the creep behaviour [10]. the additional free volume approach [5] allows for describing nonlinear viscoelastic behaviour under small to moderate strains. 2.1. microplane model m4 the microplane material model m4 [7] is the fourth version of the microplane model introduced by bažant [11, 12]. this constitutive model is characterized by the relationships between stresses and strains projected onto a surface of a microplane, which is oriented by its own normal vector. the basic idea is that the strain vector −→εn on the microplane is a projection of a strain tensor ε. then the normal strain on the microplane have the form εn = nij εij . (1) similarly, the shear components, which are characterized in the direction m and l, are given by orthogonal vectors, −→m and −→ l , take the form εm = mij εij ; εl = lij εij . (2) in which mij = (minj ) + (mj ni)/2 and lij = (linj + lj ni)/2 [13, 14]. all three components can be seen in fig.1(a). the static equivalence is computed approximately by the principle of virtual work written for the surface ω of a unit hemisphere 2π 3 σij δεij = ! ω (σn δεn + σlδεl + σm δεm )dω. (3) this equation represents that the virtual work of macro-stresses within a unit sphere must be equal to the virtual work of micro-stresses regarded as the tractions on the surface of the sphere [7]. substituting δεn = nij δεij , δεm = mij δεij and δεl = lij δεij , and noting that the last variational equation must 131 http://dx.doi.org/10.14311/app.2021.30.0131 http://ojs.cvut.cz/ojs/index.php/app jan vozáb, jan vorel acta polytechnica ctu proceedings x 3 x 2x 1 (a) e 2 e 1 e 3 e m e l e n e t e (b) figure 1. microplane model: (a) distribution of integration points (microplane normals) system of 21 microplanes per hemisphere; (b) microplane strain components include any variation δεij , the result is the following basic equilibrium equation [12]: σij = 3 2π ! ω sij dω ≈ 6 nm" µ=1 wµs (µ) ij , (4) with sij = σn nij + σm mij + σllij . (5) this integral is an approximation by an optimal gaussian integration formula for a spherical surface, as you can see in fig.1(b). this surface is representing a weighted sum over the microplanes. in the finite element algorithms, this integral has to be computed at each integration point of each finite element in each time step [7]. the most generalized constitutive relations on the microplane level have the forms σn (t) = f tτ =0 [εn (τ ), εm (τ ), εl(τ )] , (6) σl(t) = gtτ =0 [εn (τ ), εm (τ ), εl(τ )] , (7) σm (t) = htτ =0 [εn (τ ), εm (τ ), εl(τ )] , (8) where f , l and h are functionals of the microplane strains in time t. volumetric/deviatoric decomposition of the normal strain component [15] needs to be employed to capture the poisson ratio of polymers adequately εd = εn − εv ; εv = εkk/3, (9) where εv stands for the volumetric strain and d denotes the deviatoric component. 2.2. maxwell chain the formulation of the microplane model m4 briefly described in the previous section is extended to incorporate time-dependent behaviour of epoxies. approach similar to [10] is utilized. creep is developed at every integration point, i.e. on each microplane of every integration point. on the level of microplanes, creep is implemented as linearly viscoelastic. the behavior of each maxwell element (µ), where a spring and a dashpot are connected in series, is expressed as ε̇µ = σ̇µ eµ + σµ ηµ , (10) where eµ is the elastic modulus and ηµ is the viscosity of the element. in the proposed model, the viscosity is assumed only for the deviatoric part of the normal strain component on each microplane separately. thus, according to the approximation by the generalized maxwell model, the equation for calculation of deviatoric stress predictor in the incremental form reads σved = σ i d + e ′′ d (εd − ε i d ) − ∆σ ′′ d = = σid + e ′′ d (∆εd − ∆ε ′′ d ), (11) where i means the initial value in the time step ∆t and ve stands for the viscoelastic stress increment, which is the generalization of the elastic stress increment. variables labeled by ′′ are calculated according to the equations e′′d = n" µ=1 1 − e−∆t/τµ ∆t/τµ ed,µ, (12) ∆σ′′d = n" µ=1 (1 − e−∆t/τµ )σid,µ, (13) ∆ε′′d = ∆σ ′′ d /e ′′ d , (14) where n is the number of maxwell elements and τµ = ηµ/ed,µ relaxation time of the µ-th element. if a constant poisson’s ratio ν is assumed ed,µ = 5eµ (2 + 3γ)(1 + ν) , (15) where γ is a parameter that may be chosen or can be optimized so as to match the given test data best (this value is set to 1 in this study) [16]. note that remaining constitutive relations remain unchanged and the moduli are updated based on e′′ = e′′d (2+3γ)(1+ν). 2.3. free volume the last extension of our model is a free volume approach to extend the viscoelastic behaviour to a nonlinear scale, which is more convenient for the material. the description of this approach can be found 132 vol. 30/2021 modelling of particulate thermoset polymers figure 2. reological model of the maxwell chain. τµ eµ [mpa] 1 10 2.6174 2 102 107.5458 3 103 455.0239 4 104 933.0051 5 105 562.4520 6 106 273.8973 7 107 198.4917 8 108 135.9774 9 109 41.7351 10 1010 4.0175 11 1011 4.7209 12 1012 -0.0819 13 ∞ 39.9691 table 1. used values of maxwell chain. values converted from [5, tables 1 and 2]. in [5]. in the employed approach, the dilation is introduced into the time scale, which results in a nonlinear system of equations. note that for the case of infinitesimal dilatation, the nonlinear free volume theory reverts to linear viscoelasticity. in general, temperature (t ), moisture content (c) and mechanical dilatation (θ) influence the time scale of the material in a similar manner such that a = a(t, c, θ). (16) the shift factor a is connected to the free volume as suggested in [17, 18] log10 a = b 2.303 # 1 f − 1 f0 $ , (17) where f is fractional free volume, f0 represent the fractional free volume at a reference state. the coupling of the free volume to material parameters that affect its macroscopic volume can be expressed using a linear dependence f = f0 + αv ∆t + βv ∆c + δθ, (18) where αv is the volumetric thermal expansion of the free volume, βv represent the volumetric expansion due to a change in moisture content and δ relates the change if the free volume due to mechanical volume changes of the polymer. combining eq. (17) with eq. (18) we get log a = −b 2.303f0 # αv ∆t + βv ∆c + δθ f0 + αv ∆t + βv ∆c + δθ $ . (19) this equation includes the influence of all essential properties related to the development of the volume changes. the influence of the shift factor on the stress response is through convolution, or superposition, integrals relating the time dependent stress response to the strain history [5] by σd = ! t 0 e′′d (t ′ − τ ′) ∂εd ∂τ dτ, (20) where the kernel takes the form t′ − τ ′ = ! t τ dξ a[t (ξ), c(ξ), θξ] (21) to account for the temperature, moisture and dilatation histories. note that in [5], the modified free volume approach is utilized to simulate shear-dominated loading scenarios with little or no dilatation. such an approach is not needed in the current formulation since the creep behaviour is defined on microplanes. parameter value b 0.5 f0 0.1 δ 1 αv 2.857 × 10−3 °c−1 t0 55 °c table 2. used material parameters. values taken from [5]. 3. results in this phase of the development process, we used a single-point calculation to verify the numerical model using [19]. the values for the maxwell chain from table 1 and the material parameters from table 2 were used in the simulation. the element was loaded by pure shear deformation. fig. 3 shows the force vs. displacement diagram. the blue line represents the elastic loading of the element, and it represents the upper bound for the presented results. the red line shows the elastic loading of the material, including the effect of the maxwell chain. as expected, creep affects the magnitude of the reaction, and therefore the curve has a smaller slope. the last green line represents the investigated microplane model. initially, it has the same inclination as the elastic model with the maxwell chain, but as the deformation increases, the response deviates. note that for this study, the microplane model behaviour is assumed to be purely elastic. 133 jan vozáb, jan vorel acta polytechnica ctu proceedings figure 3. results from finite element analysis: cube with an edge 1 cm, loading pure shear. 4. conclusions this paper presented a new model that can be used to describe the nonlinear creep of thermosets, which is composed of three parts: (i) the microplane model m4, which calculates elastoplasticity and damage; (ii) the maxwell chain model, which adds the effect of linear creep to the calculation; (iii) the free volume model, which extends creep to a nonlinear scale. in fig.3, we can see that according to preliminary results, the model behaves as expected because the effect of creep and softening is noticeable under shear load. at the same time, it is necessary to test the model on a larger scale at different load cases to ensure its accuracy. the arcan experiment presented in [5] would be appropriate. furthermore, it is necessary to test the model under a combination of different temperatures and humidity for verification of the free volume model parameters. if the accuracy of the model is still not sufficient, it could be caused by curing of the material which could be additionally connected to our numerical model. acknowledgements the financial support provided by the gačr grant no. 19-15666s and by the czech technical university in prague within sgs project with the application registered under the no. sgs20/155/ohk1/3t/11 is gratefully acknowledged. references [1] m. a. meyers, k. k. chawla. mechanical behavior of materials. cambridge university press, 2nd edn., 2008. [2] j.-p. pascault, h. sautereau, j. verdu, r. j. williams. thermosetting polymers, vol. 64. crc press, 2002. [3] r. f. landel, l. e. nielsen. mechanical properties of polymers and composites. crc press, 1993. [4] p. silva, t. valente, m. azenha, et al. viscoelastic response of an epoxy adhesive for construction since its early ages: experiments and modelling. composites part b: engineering 116:266–277, 2017. [5] c. f. popelar, k. m. liechti. multiaxial nonlinear viscoelastic characterization and modeling of a structural adhesive. journal of engineering materials and technology 119(3):205–210, 1997. https://asmedigitalcollection.asme.org/ materialstechnology/article-pdf/119/3/205/ 5681124/205_1.pdf doi:10.1115/1.2812245. [6] g. cusatis, d. pelessone, a. mencarelli. lattice discrete particle model (ldpm) for failure behavior of concrete. i: theory. cement and concrete composites 33(9):881–890, 2011. [7] z. p. bažant, f. c. caner, i. carol, et al. microplane model m4 for concrete. i: formulation with work-conjugate deviatoric stress. journal of engineering mechanics 126(9):944–953, 2000. [8] f. c. caner, z. p. bažant. microplane model m4 for concrete. ii: algorithm and calibration. journal of engineering mechanics 126(9):954–961, 2000. [9] g. di luzio. a symmetric over-nonlocal microplane model m4 for fracture in concrete. international journal of solids and structures 44(13):4418–4441, 2007. [10] z. p. bažant, f. c. caner, m. d. adley, s. a. akers. fracturing rate effect and creep in microplane model for dynamics. journal of engineering mechanics 126(9):962–970, 2000. [11] z. bažant. microplane model for strain-controlled inelastic behaviour, chapter 3. mechanics of engineering materials, chapter 3 1984. [12] z. p. bažant, p. g. gambarova. crack shear in concrete: crack band microplane model. journal of structural engineering 110(9):2015–2035, 1984. [13] z. p. bažant, p. c. prat. microplane model for brittle-plastic material: i. theory. journal of engineering mechanics 114(10):1672–1688, 1988. [14] z. p. bažant, p. c. prat. microplane model for brittle-plastic material: ii. verification. journal of engineering mechanics 114(10):1689–1702, 1988. [15] i. carol, z. bažant. damage and plasticity in microplane theory. international journal of solids and structures 34(29):3807–3835, 1997. [16] z. bažant, y. xiang, p. prat. microplane model for concrete. i: stress-strain boundaries and finite strain. journal of engineering mechanics 122(3):245–254, 1996. [17] a. k. doolittle. studies in newtonian flow. ii. the dependence of the viscosity of liquids on freeâăřspace. journal of applied physics 22(12):1471–1475, 1951. https://doi.org/10.1063/1.1699894 doi:10.1063/1.1699894. [18] a. k. doolittle, d. b. doolittle. studies in newtonian flow. v. further verification of the freeâăřspace viscosity equation. journal of applied physics 28(8):901–905, 1957. https://doi.org/10.1063/1.1722884 doi:10.1063/1.1722884. [19] es3. mars modeling and analysis of the response structures. 134 https://asmedigitalcollection.asme.org/materialstechnology/article-pdf/119/3/205/5681124/205_1.pdf https://doi.org/10.1115/1.2812245 https://doi.org/10.1063/1.1699894 https://doi.org/10.1063/1.1699894 https://doi.org/10.1063/1.1722884 https://doi.org/10.1063/1.1722884 acta polytechnica doi:10.14311/app.2020.27.0048 acta polytechnica 27(0):48–52, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app comparison of mo coating deposited by twas and aps sinplexpro thermal spray techniques šárka houdkováa, ∗, jakub antoša, petra šulcováa, kateřina lencováa, jan hajšmanb, karolína burdováb, josef duliškoviča a výzkumný a zkušební ústav plzeň, tylova 1581/46, 301 00 plzeň, czech republic b západočeská univerzita v plzni, fakulta strojní, univerzitní 2732/8, 301 00 plzeň, czech republic ∗ corresponding author: houdkova@vzuplzen.cz abstract. the wear resistant mo coatings were deposited using smartarc electric arc wire spray system and advanced sinplexpro high throughput atmospheric plasma spray gun. the coatings were analyzed in terms of microstructure, mechanical properties and tribological characteristics, including wear rate and wear mechanism. it was shown that despite of the differences in deposition process principles, the properties of the coatings don’t differ with respect to their functional properties. both coatings reached comparable hardness, microhardness, wear resistance and coefficient of friction. the economic aspects remain the main argument for the recommendation of the most suitable thermal spray deposition technique for the application of molybdenum coatings. keywords: aps, astm g-133, coating, cof, mo, nanoindentation, sliding thermal spray, twas, wear. 1. introduction thermally sprayed mo coatings are widely used as surface treatment offering high wear and scuffing resistance. due to their superior tribological behavior they are applied especially in sliding contact [1, 2]. to deposit the mo-based coating, various thermal spray techniques can be utilized. as the coating’s microstructure is dependent on the used thermal spray method the differences in the wear resistance and friction behavior can be also expected [3]. the atmospheric plasma spraying (aps) is the most widespread technique for thick coating deposition providing the possibility to spray a wide range of materials, including metals, hardmetals or ceramics with various functionalities, it is also usually utilized to deposit the wear and scuffing resistant mo coatings. in addition to the examination of coatings sliding properties [2, 4, 5], the spraying of mo was also used for studying the process-structure-property relationships and for creating so called "process maps" of thermal spraying [6, 7]. as various studies were previously published on available aps systems, there is a lack of information concerning the recently developed sinplexpro high throughput atmospheric plasma spray gun. the twin wire electric arc spraying (twas) is a technique utilizing the material in the form of wires. the principle of the twas limited its use on the materials that are electrically conductive. that is why it is used primarily for spraying of fe and ni-based alloys, even though the wires with more complex compositions are also available [8]. the lower operating costs, as well as lower price of feedstock material in the form of wire compared to powder, can make the twas technique desirable alternative to plasma spraying if it produces the mo coatings with sufficient properties. in the paper, the comparison of mo coatings, sprayed by aps sinplexpro plasma torch and smartarc electric arc wire spray system is done to evaluate the differences in the coatings microstructure, properties and deposition efficiency. the aim is to recommend the most suitable spraying system for the application of mo coatings on the sliding surface of piston rings for the automotive industry. 2. experimental feedstock material: two kinds of feedstock material were used: (i) the pure (99.55%+) molybdenum powder amdry 313x from oerlicon metco, agglomerated and densified, with (45-75 µm) size range for aps spraying and (ii) the pure molybdenum (99.9%+) wire w400.1 from flame spray technologies, 1.6 mm diameter. spraying: the sinplexprotm plasma torch and the smartarctm twas spraying system, both from oerlicon metco, were used for spraying onto grid blasted (al2o3, 0.8-1 mm grain size) carbon steel substrates (45x25x5 mm). the spraying parameters were previously optimized for reaching the highest coating quality. to evaluate the deposition efficiency, the consumption of feedstock material was measured during spaying of 1 mm thick coating onto the prototype part piston ring spine ∅ 120 mm, 280 mm long. microstructure evaluation: the scanning electron microscope (sem) evo ma25 from zeiss with lab6 thermal filament and equipped by edx detector sdd 48 http://dx.doi.org/10.14311/app.2020.27.0048 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 mo coating: twas vs. aps sinplexpro figure 1. the sem of mo coatings sprayed by a) aps and b) twas technique figure 2. the coatings microhardness depth profile x-max 20 oxford instruments was used to evaluate the microstructure of the coating on the crosssections. the phase compositions were analysed by x-ray diffraction (xrd), using d8 discover diffractometer with 1d detector and cokα radiation. the obtained diffraction patterns were subjected to quantitative rietveld analysis performed in topas 4.2 [9] in ipp ascr. coating properties evaluation: the adhesion tests were realized in accordance with astm c633-13 [10]. for each value, at least 3 measurements were done. surface roughness was measured in accordance with en iso 4288 [11]. for each value, at least 3 measurements were done. the surface hardness hr 15n was measured on the gently ground surface of coatings. for each value, at least 5 measurements were done. microhardness hv0.3 was measured in the middle of coatings cross-sections. for each value, at least 7 measurements were done. microhardness depth profile hv0.1 was measured on the coating’s cross sections in 0.05 µm steps. for each value, at least 3 measurements were carried out in one depth. the instrumented hardness was measured in rti uwb using the nanotest vantage testing device from micro materials company, equipped with berkovich indenter, by 5 n testing load, 200 mns−1 loading rate, 10 s dwell period at maximum load, 20 indents were performed 10 µm below the surface and 20 indents 100 µm above the interface. the sliding wear resistance was measured by ball-onflat test, in accordance with astm g-133 [12], using 6 mm diameter. cr-steel ball counterpart; 25 n load, 50 hz frequency; 1000 s test duration. during the test, the coefficient of friction was recorded. the coatings sliding wear resistance k [mm3/nm] was determined from the wear tracks profile measurements. the average value of three astm g-133 tests on each sample is reported. 3. results and discussion the microstructure of the aps and twas sprayed coatings is shown in figure 1. in both cases, the porosity, as well as cracks of the individual particles (splats), can be observed. the size of the splats is lower in the case of aps coating, resulting from the fine powder compared to the splats originated from the molten wire tips. both coatings have a variable microstructure across the cross-sections. less porosity and inter-splat decohesion was observed in the lower part of the coatings, compared to the upper part. this phenomenon is still unclear, nevertheless, its intensity is dependent on the spraying parameters of each technology, related 49 š. houdková, j. antoš, p. šulcová et al. acta polytechnica figure 3. coefficient of friction in dependence on the time of the test for a) aps and b) twas sprayed coatings mo coating hit (gpa) eit (gpa) hit (gpa) eit (gpa) upper layer upper layer lower layer lower layer aps 3.76 ± 0.43 138.6 ± 8.6 3.39 ± 0.33 103.7 ± 10.0 twas 3.31 ± 0.69 83.3 ± 8.4 2.98 ± 0.46 66.4 ± 8.5 table 1. results of instrumented indentation. probably to the temperature and velocity of the particles, as well as to intensity and frequency of cooling during spraying. the amount of energy, transferred to the coated sample can be responsible for coatings densification. simultaneously, the influence of transferred heat and impact energy can be connected with changes of substrate hv0.1 microhardness (see fig. 2). the phase composition was evaluated by xrd. both coatings composed of mo, accompanied by a small amount of moo2. according to the rietveld quantitative analyses, its amount reached 0.76 wt.% for aps coating and 2.44 wt.% for twas coating. the thicknesses of the coatings were measured 0.48 ± 0.02 mm and 0.77 ± 0.01 mm for aps and twas coating resp. as the number of passes necessary to reach the thickness was 10 and 6, the thickness per pass was 48 µm and 128 µm for aps and twas coating resp. the consumption of feedstock material was measured during spraying the piston ring spine (∅120 mm-280 mm long). to spray 1 mm thick coating, the 0.2 kg of powder was consumed during aps spraying, while 1.83 kg of wire was needed for twas spraying. the adhesion of the coatings was 37.4 ± 5 mpa for aps and 32 ± 5 mpa for twas coating. these values are lower than values, found in ref [3]. coating surface roughness was 8.62 ± 0.27 µm ra and 56.85 ± 3.80 µm rz for aps and 18.89 ± 0.92 µm ra and 114.52 ± 4.01 µm rz for twas coating. while the roughness of aps coating is in agreement with ref. [3], the ra of twas coating is much higher in [3] (18.89 vs. 9.81 ra). the surface roughness is connected with the spreading of splats after the droplets impact and can be considered as a parameter of coatings quality. surface hardness hr 15n reached comparable values (67.4 ± 1.4 and 67.8 ± 1.5 for aps and twas coating resp.). the surface hardness value involved besides the property of the material, also the contribution from present pores, cracks and intersplat boundaries. the microhardness hv0.3 of the coatings, measured on the coating’s cross-sections was slightly lower for aps coating compared to twas (327 ± 20 and 387 ± 37 for aps and twas coating resp). the presence of a higher amount of hard oxide inclusions can be responsible for higher microhardness of twas coating. the overall hv0.3 values are in agreement with the microhardness measured in ref. [1, 3, 4]. to evaluate the variation of coating properties across the cross-sections, the microhardness depth profile was done. as can be seen from figure 2, there is not a significant difference between aps and twas coating. in both cases, the hardness increases towards the surface. such observation seems to be in contradiction with the assumption, that the less porous microstructure close to the coating-substrate boundary will be harder than the porous layer in the upper part of the coating. to confirm the microhardness results, the instrumented indentation measurement was done on both coatings in the upper porous and lower dense layer of each coating. the load of 5 n was used for indentation to include not only the inner splats properties but also the influence of porosity and intersplats cohesive strength. the measured values of hit and eit are summarized in the table 1. even though high scatter of the measured values given by heterogeneous microstructures of both coat50 vol. 27/2020 mo coating: twas vs. aps sinplexpro figure 4. wear mechanism for aps (a,b,c) and twas (d.e.f) sprayed coatings ings, it can be observed that the lower dense close-tointerface layer has a lower hardness and lower modulus of elasticity, compared to a higher, more porous layer close to surface. this observation is in agreement with the hv0.1 depth profile measurement. during the ball-on-flat linear oscillating test, the coefficient of friction (cof) was recorded and can be seen in figure 3. similar behavior of both types of coatings can be seen. the continuous slow increase was followed by a sudden drop of cof value in 400600 s of the test. after, the constant cof value of 0.7 was recorded till the end of the test. the sudden decrease of cof is usually connected with the creation of stable tribolayer in the wear track. also, the transfer of worn coating material onto the surface of the counterpart can be another reason for cof decrease, however, it was not analyzed during this study. the wear resistance of both types of coatings is very similar. for aps sprayed coating, wear coefficient k was 1.91*10-4 ± 1.9*10-5 mm3/nm, while for twas coating 2.06*10-4 ± 2.8*10-5 mm3/nm. for comparison, the wear coefficient of aisi 316l steel is 2.74 ± 1.37. in [2] the reported wear rate was much lower compared to the results of samples evaluated in this study. the differences in wear tests (astm g-99 vs astm g-133) and particularly in used loads (5n vs 25n) can play a significant role in the wear rate. the wear mechanism can be seen in figure 4. for both types of coatings, the delamination of the parts or whole splats was identified (fig. 4b,e) as the main mechanism responsible for wear. except of delam51 š. houdková, j. antoš, p. šulcová et al. acta polytechnica ination, ploughing takes place. in the wear track, the tribo-layer was locally identified (fig. 4c,f). the tribo-layer usually consist of deformed coating wear debris, transferred counterpart material and oxides, originated during sliding [13]. the edx analyses show the presence of fe (ca. 3-4 wt.%) and cr (ca. 1-1.5 wt.%) in the wear track, originated from the counterpart, together with 11-14 wt.% of oxygen confirming the assumption of tribo-oxidation. there are no differences in the wear mechanism between aps and twas sprayed coating, except the size of delaminated areas. they are bigger in the case of twas coating, which is connected with bigger splats dimensions (see fig. 1). the bigger size of delaminated splats is than responsible for slightly higher wear volume. 4. conclusions despite of the differences in deposition process principles, the properties of the coatings don’t differ with respect to their functional properties. both coatings reached comparable hardness, microhardness, wear resistance and coefficient of friction. wear mechanisms also don’t differ splat delamination and ploughing are the main observed mechanism responsible for wear. in the case of twas, the size of delaminated areas is bigger in consequence of the bigger size of original splats. considering the economic aspects, the significantly lower deposition efficiency of twas technology and longer time of deposition makes the twas 70 % more expensive than aps sinplexpro for spraying of a reference component, despite of more than 2 times higher unit price of mo powder compared to wire. acknowledgements the paper has originated in the framework of the project of technology agency of the czech republic no. tj02000135. references [1] j. ahn, b. hwang, s. lee. improvement of wear resistance of plasma-sprayed molybdenum blend coatings. journal of thermal spray technology 14(2):251–257, 2005. doi:10.1361/10599630523827. 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[13] g. bolelli, b. bonferroni, j. laurila, et al. micromechanical properties and sliding wear behaviour of hvof-sprayed fe-based alloy coatings. wear 276-277:29 – 47, 2012. doi:10.1016/j.wear.2011.12.001. 52 https://doi.org/10.1361/10599630523827 https://doi.org/10.1016/j.wear.2007.01.018 https://doi.org/10.1007/s11666-018-0706-2 https://doi.org/10.1016/s0043-1648(98)00395-0 https://doi.org/10.1016/s0043-1648(01)00842-0 https://doi.org/10.1016/s0921-5093(02)00642-1 https://doi.org/10.1016/j.msea.2005.04.056 https://doi.org/10.1107/s0021889869006558 https://doi.org/10.1016/j.wear.2011.12.001 acta polytechnica 27(0):48–52, 2020 1 introduction 2 experimental 3 results and discussion 4 conclusions acknowledgements references doi:10.14311/app.2021.30.0036 acta polytechnica ctu proceedings 30:36–40, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app simulation of t-joints between rhs steel members with offset in abaqus cae svitlana kalmykova czech technical university in prague, faculty of civil engineering, department of steel and timber structures, czech technical university in prague, thákurova 7, 166 29 prague 6, czech republic correspondence: svitlana.kalmykova@fsv.cvut.cz abstract. the current paper focuses on numerical simulation peculiarities of offset welded rectangular hollow section joints. understanding the modelling techniques can result in easier and faster and above all correct outcomes from fea for future use. the steel joints under discussion are composed from cold-formed regular rectangular hollow sections where rhs brace members are laterally shifted from chord axis. joints work under monotonically increasing compression load applied to a brace top. numerical models were developed in fe programme abaqus. fe-models is composed of c3d8r 8-noded solid linear brick elements with an emphasis on mesh size effect and modelling of a weld seam. fe advanced model were compiled considering both material and geometric nonlinearities. for validation purposes, the full-scale laboratory tests were conducted. proposed fe models reliably predict the structural behaviour of welded offset t-joints thanks to good agreement achieved on deformation limit 3 % b0 with the maximum deviation 10.3 %. keywords: fe programme abaqus, hollow structural sections (hss), numerical modelling, offset t-joints, rectangular hollow sections (rhs), truss connections. 1. introduction steel structural hollow sections, including rectangular hollow sections (rhs), are nowadays one of the most efficient load-bearing elements of the lattice structures. despite its wide-spread application, there is still lack of design techniques for joints produced with offsets. since the mid-1960s, vast study on hss connections structural behaviour has been undertaken. the results were implemented in current regulations and standards such as en 1993-1-8 (2005), ansi/aisc 360-05 (2005), csa-s16-09 (2009), iso 14346:2013 or a design guide cidect dg3. it is essential to note that existing methods of hss connections design are for coaxial ones only, i.e. centreline of a branch member should have an intersection with centreline of a chord. the term "standard" is used for geometry of aforesaid connections. at the same time, the question concerning mechanical resistance of joints under discussion is practically uncovered by design recommendations and guidelines. as of today, there are few studies published on the topic, namely, the earliest experimental research of laterally offset tand n-joints done by dawe et al. [1], kalmykova and wald [2, 3] including laboratory and numerical research, and wei and packer [4] who performed a set of the offset shs tand xjoints laboratory testing and proposed certain design approaches simultaneously with an analytical equation to predict resistance of the offset joints. the strength, modes of failure and loaddisplacement curves can be gained from a proper numerical analysis. considering costliness of physical laboratory experiments, it is of high impotence to find ways of effective numerical simulation of joints (see figure 1) to trustworthy predict their structural behaviour. figure 1. welded offset t-joint geometry. the aim of the current study is to present a practical numerical simulation guideline for offset joints and connections using abaqus cae. 2. finite element modellings 2.1. general numerical modelling of the fully welded offset tjoints was done in abaqus. it is general-purpose finite element programme, release 6.14. fe models of joints were composed in compliance with pren 36 http://dx.doi.org/10.14311/app.2021.30.0036 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 rhs steel members t-joints simulation figure 2. finite element mesh of t-joint. 1993-1-14: 2020. static, riks procedure available in the abaqus library was employed to run loaddisplacement nonlinear analysis. fe models of offset joints were built considering both geometric and material nonlinearities. carefully performed sensitivity study allowed to define the certain element type, fe mesh size of finite elements as well as welding material property. it brings accurate results with adequate computational costs. boundary conditions, loading, weld seam modelling and simulation of materials are fundamental inputs in performed fe analysis. 2.2. geometry of joints experimental programme includes two types of tjoints employed for the validation procedure. the first type presents a regular coaxial geometry of tjoint without any offset; the second one was produced with the brace lateral offset (see figure 1). both joints were loaded by the monotonically increased axial compression applied at the brace top. brace member is made of cold-formed rectangular hollow section rhs 50 × 30 × 4 mm en10219. coldformed rhs 150 × 100 × 4 mm en10219 is used for chord member. 2.3. element types and fe mesh size in a brick model the element type used for joint modelling is 3d 8node solid element with reduces integration (c3d8r) from abaqus library. it is fully isoperimetric element with the first order interpolation. finite element mesh of a model is presented in figure 2. it consists of two regions with two various fe element sizes. smaller mesh region comprises fe bricks with the sizes up to 4 × 4 × 4 mm and is distributed along the direct chord axis on the distance of 240 mm. the distance of smaller mesh region for the brace member along the direct brace axis equals 300 mm. the rest of the fe mesh consists of 8 × 8 × 8 mm elements for both members of the joint. additionally, at the corner areas, the finer 4-element mesh was employed because of great importance to transfer stresses from a flange to a web of a joint member. isoparametric angles were monitored as there was no distortions in the fe bricks. aspect ratio of the fe bricks were limited to be less than 3. the number of in-depth fe layers for chord and brace walls were monitored in order to analyse model sensitivity. the number of layers were varied from 1 to 6. based on the obtained data and subsequent analysis, 3-layer mesh was chosen for further simulations by virtue of ability to provide accurate results with reasonable computational costs. apart from c3d8r brick fe units, two more types of solid 3d elements were utilized. there are c3d8i and c3d20r. it is to check processing time and precision of obtained results for fe models made by it. 2.4. loading and weld simulation joints work under monotonically increasing compression load applied at a brace top. axial compression was imposed to the reference point on an analytical rigid plate (see figure 3). weld seam, in general, has considerable impact to the mechanical resistance of t-joints. this is particularly the case for the fillet weld corner areas. thus, weld was employed in fe models of offset joints. as for the brick model, weld seam is the only member that joining brace and chord. thereto, a yawns or gaps were modelled between external face of the chord top flange and bottom section of the brace with the help of advanced simulation on initial geometry creation step. specific procedure was applied to the meshing. as for the shell model, one should go to the section 2.7. for simplicity, the material properties of the weld seam and steel members were assumed the same. it is due to complexity of weld material property investigation when an electrode’s composition is mixed with the main material of a joint. fe solid elements types coinside with the chosen type for a model. 2.5. boundary conditions as for boundary conditions of the joint, the chord member was stacked on the rigid base, i.e. it is fully restrained from rotation and displacement. additionally, the displacement in the direction of the applied 37 svitlana kalmykova acta polytechnica ctu proceedings figure 3. displacement of the reference point and boundary conditions on the model. load is allowed for the top brace edge, while other displacements and rotations are prohibited. 2.6. material tensile coupon tests (see figure 4) were performed in order to provide real material properties to an advanced fe model. the coupons were cut in the rolling direction from additional pieces of rhs made from the same material. a total of 5 tensile tests were conducted: 3 tests on s355 j2h material taken from a chord element section, and 2 tests on s355 j2h taken from a brace element profile. the middle part of a profile face was chosen for cutting coupons to avoid zones affected by hardening. material properties including young’s modulus, the ultimate tensile stress and the corresponding strain were recorded. more detailed information including measured data and true stress-strain curves is provided in kalmykova and wald [2]. in general, for elastic-plastic bodies analysed by abaqus, material properties of a specimen are assigned in two dialog boxes, namely, mechanical elastic and plastic material behaviour dialog boxes. the elastic response of steel is modelled by linear elasticity with young modulus. elastic properties are specified as isotropic (see table 1). to model isotropic plasticity in abaqus, the mises yield surface is used. the plasticity data are determined by true stress versus logarithmic plastic strain (see table 2). fracture strain stress triaxiality strain rate 1 0.3 0 0.3 1 0 table 1. material behavior assigned in abaqus (elastic isotropic). for the specimen modelling in abaqus, the real material model resulting from the coupon tensile tests was used. the initial part illustrates proportional yield stress plastic strain 320 0 355.594 2e-005 365.568 0.01806 420 0.04679 459.563 0.07013 492.25 0.09297 517.5 0.11532 537.625 0.1372 567 0.17962 587.5 0.22035 table 2. material behavior assigned in abaqus (plastic isotropic). dependency of stress and strain, an evidence of elastic property, up to the tensile yield strength fy with recorded young’s modulus e and poisson’s ratio ν. beyond the elastic limit, ultimate tensile strength fu and elongation after fracture εf measured on coupon tests create the post-yield response of the bilinear material model. additionally, von-mises yield rule and kinematic hardening model were employed into the model. extensive plastic strain is involved into nonlinear fe analysis. for that reason, static stress-strain diagram measured on coupon tests was converted to true stress-logarithmic plastic strain diagram. abaqus user’s manual contains recommendations based on which the true stress σtrue and logarithmic plastic strain εpl were computed. during the simulation, an attempt to simulate the elimination from calculations of individual physically damaged finite elements was made since the strength criterion for such elements is no longer satisfied. damaged cells were excluded from the calculation by introducing the ductile damage option (see table 3). 38 vol. 30/2021 rhs steel members t-joints simulation figure 4. coupon tests specimens and scheme of loading. young’s modulus poisson’s ratio 201000 0.3 table 3. material behavior assigned in abaqus (ductile damage). 2.7. shell fe model in addition to the main procedure of numerical model validation, it was decided to conduct comparative analysis of similar fe-models composed by solid elements (c3d8r) and their counterparts constructed by shell finite elements. the element type used for joint modelling is a 4node doubly curved shell element with reduced integration (s4r) from abaqus library. a 5-point integration was utilized by the thickness of the shell with all degrees of freedom per node. in modelling, sections of the joints’ brace and chord were presented by their middle surfaces with assigned thicknesses which are correspondent to the real thicknesses of components. as it was stated in more earlier studies (korol and mirza [5], matos et al. [6]), shell finite elements are able to assure precise computation to overwhelming majority of applications owing to transverse shear deformation presence that crucial for thick shell elements simulation (see figure 5). a four-node doubly curved shell element with reduced integration (s4r) has been used to model brace and chord members. a five-point integration was applied through the shell thickness with a full complement of six degrees of freedom per node. in modelling, sections of the joints’ brace and chord were presented by their middle surfaces with thicknesses which are correspondent to the real thicknesses of components. as it was stated in more earlier studies (korol and mirza [5], matos et al. [6]), shell elements provide an accurate solution to most applications by allowing for transverse shear deformation, which is important for simulation of thick shell element figure 5. as a simplification, some of researchers omit the weld seam in shell fe models of welded hss joints. this usually creates undervaluing of the joint strength. nevertheless, for the specimens under discussion, the weld seam was considered. filled weld with throat thickness of 10 mm at the intersection of brace and chord was modelled according to van der vegte et al. [7]. inclined at an 45 angle circle of elements is displaced along the brace and chord intersection area. some of the researchers omit the welds for simplicity in the modelling of welded tubular joints using shell elements, which leads to underestimating the joint strength. however, for the specimens under discussion, the weld seam was considered. filled weld with throat thickness of 10 mm at the intersection of brace and chord was modelled according to van der vegte et al. [7] using a "ring" of shell element along the brace and chord intersection region. 3. experimental study full-scale laboratory experiments were carried out by kalmykova and wald [2] with the aim to study failure modes of the standard and offset t-joints specimens and to get data for the validation of fe models developed for numerical studies. 4. validation for the purposes of this paper, results from numerical fe advanced model in abaqus were calibrated against relevant full-scale experimental data for both 39 svitlana kalmykova acta polytechnica ctu proceedings figure 5. "shell" fe models weld seam. standard and laterally offset t-joints. abaqus fe models were validated against experimental results conducted by the authors mentioned hereinabove. validation procedure and its complete outcomes have been provided in detail in [3]. the comparison of the "load-displacement" diagrams coming from the fullscale experiments and fe analysis is shown on figure 6 and figure 7. figure 6. validation of fe models against laboratory tests for a standard joint. test and finite element analysis results show good agreement achieved for both standard and offset models on allowed deformation level of 0.03 b0 with the maximum difference of 10.3 % and the minimum difference of 2.6 %. 5. conclusion fe models are an essential tool for structural behaviour assessment of steel offset joints between hss members. the structural response predicted by these numerical models adequately reflects the actual behaviour of the joints. to guarantee it, the fe models were validated on the results of physical laboratory tests. the fe analyses results accurately aligned with the test outcomes. hence, it could be said that the proposed fe models reliably predict the structural behaviour of welded offset t-joints with rhs members. 3d 8-node solid elements with reduces integration (c3d8r) is recommended for further simulation of offset t-joints between rhs members as the ones figure 7. validation of fe models against laboratory tests for an offset joint. that provide accurate results in adequate computation time. 6. list of symbols e young’s modulus ν poisson’s ratio σu ultimate tensile stress εu ultimate strain fy tensile yield stress fu ultimate tensile strength σtrue true stress εpl logarithmic plastic strain 7. acknowledgements the research was carried out under the grant sgs19/150/ohk1/3t/11. references [1] j. l. dawe, y. liu, r. zhang, a. dukuze. experimental study of offset hss connections. journal of structural engineering, american society of civil engineers 132(5):717–725, 2006. [2] s. kalmykova, f. wald. full-scale experimental study of welded rhs t-joints with offset and its main outcomes. grant journal 8(1):76–81, 2019. [3] s. kalmykova, f. wald. welded t-joints between rectangular hollow section members with offset. in proceedings of xii conference on steel and composite construction 1:1157–1167. [4] f. wei, j. a. packer. experimental study of laterally offset rhs x-connections. [5] r. m. korol, f. a. mirza. finite element analysis of rhs t-joints. journal of the structural division 108(9):2081–2098, 1982. [6] r. m. m. p. matos, l. f. costa-neves, l. r. o. lima, et al. resistance and elastic stiffness of rhs "t" joints: part i axial brace loading. latin american journal of solids and structures 12:2159–2179, 2015. [7] g. j. van der vegte, c. h. m. de koning, puthli, et al. numerical simulation of experiments on multiplanar tubular steel x-joints. international journal of offshore and polar engineering 1(3), 1991. 40 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2021.31.0030 acta polytechnica ctu proceedings 31:30–35, 2021 © 2021 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague calculation of the minimal length of the high-speed line petr nachtigall∗, erik tischer university of pardubice, jan perner transport faculty, department of transport technology and control, studentská 95, 532 10 pardubice, czech republic ∗ corresponding author: petr.nachtigall@upce.cz abstract. this paper explores the minimum length of continuous sections of the high-speed line. the minimum length is examined in terms of the maximum speed of high-speed vehicles. both traditional trainsets consisting of traction units and coaches, and train units were selected for examination. the graphs present the difference in the ability of the different vehicles to reach and use the maximum speed. keywords: high-speed line, opentrack, simulation, track speed. 1. introduction the construction of a high-speed network is going to be a reality in the czech republic. after a long time of uncertainty, there is a real plan of activities that will make us part of the map of a high-speed railway. as the construction of the entire network will take a long time, this paper is trying to calculate the minimum length of a continuous line to make the high-speed traffic attractive for passengers. the opentrack simulation tool was used for simulation [1]. subsequent calculations were performed in ms excel. to find out which track length is effective for the defined maximum track speed, we analyzed rides of predefined trainsets on a hypothetical infrastructure. the history of high-speed transport dates to the 20th century and has experienced its greatest upswing since the beginning of the 21st century. the total length of high-speed lines has already exceeded 50,000 km and it is planned to reach the target of 100,000 km in the next 15 years. as for individual countries, the most progressive country is china with its more than 35,000 km, with a view to reaching 70,000 km by 2035. europe lags behind in this respect with about 9,000 km. the czech republic is still without a high-speed railway; however, pilot projects are already being implemented and some parts of the conventional infrastructure are being built to allow a speed of 200 km·h−1. with the construction of the high-speed network, there comes the question of how long a section of the high-speed line must be to allow for the maximum track speed, making efficient use of the funds invested in this infrastructure [2, 3]. in their previous research, the authors investigated the optimum distance between two stops ensuring energy reliability. at the same time, they explored the effectiveness of introducing a speed of 200 km·h−1 on short sections of the conventional network. this paper examines the minimum length of a section of the high-speed line, which will enable the trains to reach the planned track speed. this problem is covered by the question of the efficiency of investment into the construction and operation of this specific infrastructure. there is a clear linear correlation between the track speed and construction, noise and safety measures. at the same time, a higher track speed puts stronger demands on vehicle design, power, and traction power consumption. a littler aside is the question of the minimum length of the high-speed line. this research can be used in the preparation of hsl network [4] constructions or overall preparation of the systematic timetables [5]. the czech republic is not as big as for instance china, but during the expansion of the hsl network, we can use an optimization mathematical apparatus to ensure the maximum transportability and utility of the highspeed system [6, 7]. the issue of energy efficiency is dealt with in literature [8], describing the perspective of the use of the coasting and economic modes of high-speed trains. the question of the interlocking system is answered by the etcs because the czech republic is going to install this system on the main corridors, including high-speed lines [9–14]. the costsâăof infrastructure are figuratively paid by passengers in the price of the ticket, and this price should be in accordance with the travel time and the quality of service. another important question is energy consumption. this issue is the focus of other papers [15–17]. 2. preparation of the simulation model the opentrack simulation software was used for the research. the authors have been using this tool for a long time with good results, and this software is a globally established simulation tool [18]. the model was designed with the emphasis on allowing all the trainset movements modelled to be performed in a very simple way. the simulation model was therefore 30 https://doi.org/10.14311/app.2021.31.0030 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 31/2021 calculation of the minimal length of the high-speed line prepared as simply as possible. the simulation model was created in two phases. in the first phase, an infrastructure model was created. the infrastructure model developed for this high-speed vehicle research consists of four tracks of 50 km each with no curves or slopes. in the second phase, the parameters of each trainset were entered. most important are the tractive characteristics of trainsets, maximum speed, weight, adhesive weight, and vehicle resistance. as for the research trainsets, the 151.0 engine with 7 bmz coaches, škoda 109e with 7 bmz coaches, siemens viaggio comfort non-traction unit with taurus engine and multiple units of the 680 čd "pendolino" series, ave, and ice 3 were chosen. for a more detailed description of the model, see previous article [19]. what is important for the actual simulation are the input characteristics of the individual trainsets. these characteristics are provided for in the following table 1. to calculate the rolling resistance of the individual trainsets, the davis equation was used, as expressed in equation 1. the individual parameters are included in the vehicle database within the opentrack system used for the simulation. had there been tunnels on the line, we would have had to add resistance caused by using the tunnel. however, no tunnel was included in the model. rlz = a + b ∗ ν + c ∗ ν2 [n] (1) where: rlz= train air resistance [n], ν = train speed [km·h−1], a, b, c= parameters. the simulation was running continuously. depending on the required accuracy [20], it was possible to establish the situation of the train in every moment in time. this is important in examining the impact of variables that can be easily included in our generic model, as appropriate. the actual train speed and distance covered can be calculated using integration according to equation 2 and 3. the resulting values can be used to create tachograph curves or the monitored parts thereof. ν = νp + ∫ t2 t1 a · dt [m · s−1] (2) s = sp + ∫ t2 t1 ν · dt [m] (3) where: ν = speed [m·s−1], νp= initial speed [m·s−1], t = time [s], t1 = initial time [s], t2 = target time [s], a = acceleration [m·s−2], s = distance covered [m], sp = initial distance covered [m]. all simulations were carried out under good adhesion conditions; the value of adhesion was set at 100% of the adhesion coefficient. by setting the value of adhesion utilization, it is possible to simulate potential degraded adhesion conditions. the adhesion coefficient can be calculated according to equation 4. µ = 2.1[m · s−1] ν + 12.2[m · s−1] [−] (4) where: µ = adhesion coefficient, ν = speed [m·s−1]. through the simulation, the authors intended to show the distance needed for the acceleration of each trainset. each trainset was simulated with its maximum speed. 3. results the output of the simulation is quite simple: it is a distance-speed graph for each trainset, where we can observe the basic dynamic characteristics. table 2 shows the distance necessary for a trainset to reach the speed limit. the distance required for acceleration varies greatly. it certainly depends on the tractive power of each trainset and its weight. based on this, a graph of the speed-distance dependence was created. figure 1 shows this graph illustrating the curve for each trainset. we can see that the minimum length of the highspeed line for a train to reach the limit of 200 km·h−1 lies between 4 and 10 km. if we consider a speed limit of 230 km·h−1, this distance will increase to 8 km for ice 3 and to 18 km for pendolino and railjet.âăon the other hand, those values are just minimum ones. the actual length for practical useâăshould be higher. another very important variable is acceleration. for the presentation of results, the trainsets were divided into two groups with a distance-speed graph created for each of them. shown in figure 1, the one group involves typical trainsets with a locomotive. the other group in figure 3 contains train units. it is clear that the modern railjet vehicles and locomotive 380 achieve much better results than a locomotive representing the older generation 151.0. particularly at lower speeds, both these modern locomotives have dynamic properties close to those of train units. clearly the weakest out of the train units is pendolino, which is mainly due to its older design and, therefore, its higher weight. even on a straight section, pendolino’s maximum acceleration is 0.45 m·s−2 and declines relatively quickly. starting at a distance of approximately 1 km (at approx. 100 km·h−1), its acceleration is already lower than 0.3 m·s−2. 31 petr nachtigall, erik tischer acta polytechnica ctu proceedings trainset maximum speed [km·h−1] weight [t] maximum tractive effort [kn] maximum power [kw] maximum acceleration [m·s−2] average deceleration [m·s−2] 151.0 + 7 coaches 160 443 210 4000 0.353 -0.6 škoda 109 e + 7 coaches 200 445 274 6400 0.560 -0.6 siemens viaggio comfort 230 437 300 6400 0.627 -0.6 ave s-102 300 355 200 8800 0.526 -0.6 pendolino 230 384 200 3920 0.461 -0.6 ice 3 300 463 300 8000 0.581 -0.6 table 1. basic characteristics of the vehicles simulated, source: authors. speed distance [m] [km·h−1] engine 380 engine 151.0 railjet ave pendolino ice 3 100 890 1573 737 860 970 730 160 2993 5284 2967 2610 4096 2280 200 6588 6914 4900 9400 4450 230 20804 8950 18247 7100 300 20612 330 36372 table 2. distance to be covered to reach the respective speed, source: authors. figure 1. the distance needed for acceleration [km], source: authors. 32 vol. 31/2021 calculation of the minimal length of the high-speed line figure 2. acceleration of typical trainsets [m·s−2], source: authors. figure 3. acceleration of units [m·s−2], source: authors. 33 petr nachtigall, erik tischer acta polytechnica ctu proceedings 4. conclusions this paper explores the minimum length of a section of the high-speed line for vehicles to be able to reach the maximum speed. the simulation showed that for a maximum speed of 230 km·h−1, this distance is approximately 10 to 15 kilometers. this is the distance to be covered by high-speed units to achieve this speed. an important question to be answered before the line construction is whether the line would only be used by high-speed units, or also by conventional trainsets with a maximum speed of 200 or 230 km·h−1. the simulation shows that even these trainsets can achieve their respective maximum speeds in a relatively short period of time. by contrast, older models of traction units are not suitable for operation on these newly constructed lines due to their low maximum speed and traction characteristics. encouraging is the fact that even short sections of the high-speed line are a convenient and efficient tool for the development of the railway network: not only are they parts of the high-speed network, but they can also take the strain off current conventional lines. mainly close to major urban areas, the capacity of these lines is often saturated, and they are not able to cover the ever-increasing demand for suburban transport. also freight railway undertakings are waiting for the clearing of these sections. acknowledgements this article is published within realization of the project "cooperation in applied research between the university of pardubice and companies in the field of positioning, detection and simulation technology for transport systems (positrans)", registration no.: cz.02.1.01/0.0/0.0/17_049/0008394. references [1] d. huerlimann, a. nash. opentrack simulation of railway networks version 1.3. zurich: eth zurich, institute of transport planning and systems, 2019. 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[19] p. nachtigall, e. tischer. efficiency of increasing of the track speed using simulation in opentrack. proceedings of the 32nd european modeling & simulation symposium (emss 2020), https: //www.cal-tek.eu/proceedings/i3m/2020/emss/017/ issn 2724-0029. [20] j. bulíček. cancellation of delayed trains: passengers’ and capacity points of view. žilina: matec web. conf., 2018. vol. 235. 35 https://www.cal-tek.eu/proceedings/i3m/2020/emss/017/ https://www.cal-tek.eu/proceedings/i3m/2020/emss/017/ acta polytechnica ctu proceedings 31:30–35, 2021 1 introduction 2 preparation of the simulation model 3 results 4 conclusions acknowledgements references doi:10.14311/app.2021.30.0058 acta polytechnica ctu proceedings 30:58–62, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app numerical simulation of slump flow test of cement paste composites jiří němeček∗, jiří němeček czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jiri.nemecek.1@fsv.cvut.cz (ph.d. candidate) abstract. this study presents a numerical simulation of the mini-slump flow test performed on the haegerman table, according to en 1015-3. the two-dimensional axisymmetrical volume of fluid (vof) model is used for the calculation. various scenarios with the different model set up and boundary conditions were calculated to show how the results are affected. simulations with different lifting velocities were carried out and compared to the case with instantaneous demolding, which is commonly used in numerical simulations. also, the effect on results is shown between noslip and specific shear boundary conditions on the mold. the results of simulations indicate that mold lifting should be considered in numerical modeling even if calculation time is highly prolonged. lifting velocities should not exceed the value of 0.2 m/s in order to maintain laminar flow and stable calculation. keywords: cement paste, cfd modeling, rheology, slump test, volume of fluid. 1. introduction concrete is a mixture of cement, fine, coarse aggregate, admixtures, and water. after mixing initial components with water, a fluid suitable for filling formwork is created. each mixture has different rheological properties depending on its composition, which could be characterized by various experimental tests such as slump test, l-box, v-funnel tests, viscosimeter, and vibration testing [1]. the most commonly used is the cone slump testing for its feasibility and reproducibility on in-situ and ex-situ conditions. the test is conducted with a metallic cone opened at both ends, which is placed on a metallic/wooden/glass table. the fresh mixture is filled inside the cone, and then the metallic cone is lifted in a vertical direction. when the initial shear stress of the mixture is overcome, the concrete flows by gravity [2]. the test is commonly simulated by numerical modeling using computational fluid dynamics (cfd) approach. most of the studies modeled the slump-test with instantaneous demolding [2–4]. however, such a boundary condition (b.c.) is leading to a higher slump spread. this behavior is then artificially correlated by changing constants of the material model. gao investigated the effect of a cylindrical mold lifting on the slump test of kaoline clay [5]. the results suggest the mold lifting velocity has a significant influence on the final slump spread. in this study, the setting up of a numerical model of the mini-cone slump test is shown. the effect of different boundary conditions on the final slump spread 1ph. d. candidate at ctu in prague 2associate professor at ctu in prague is investigated. 2. numerical simulation 2.1. volume of fluid model the eulerian multiphase model is the most complex model which can calculate dispersed-continuous phase interaction and continuous-continuous phase interaction. typically dispersed micrometer or millimeter size particles could be solid particles, liquids droplets, or bubbles of gas which are dissolved in a continuous phase. continuous-continuous phase interaction is represented by two immiscible phases, e.g., concrete and air. the eulerian model is solving continuity and momentum equations individually for each phase. the simplified version of the eulerian model used for continuous-continuous phase interaction is the volume of fluid (vof) model. the vof model is solving only one continuity equation and one momentum equation and tracks each phase’s volume fraction [6, 7]. the continuity equation could be simplified as ▽ · !v = 0, (1) where where !v is the velocity and time field. a single momentum equation is written as ∂ ∂t (ρ!v) + ▽ · (ρ!v!v) = − ▽ p + ▽ · ! η " ▽!v + ▽!vt #$ + ρ!g + !f , (2) where p is the static pressure, g is the gravitational acceleration, !f is the force resulting from surface tension at the interface, and ρ is the fluid density and η is the apparent viscosity. 58 http://dx.doi.org/10.14311/app.2021.30.0058 http://ojs.cvut.cz/ojs/index.php/app vol. 30/2021 slump flow test numerical simulation prog��zdoo����uljlg�erg\�prwlrq�� yhorflw\� ����������p�v d��qr�vols�e�f� e��vshfl¿f�vkhdu�e�f� d[lv�ri�v\pphwu\ fhphqw�s���� ghiruplqj�phvk dlu���ghiruplqj�phvk )j ��pp�jds e� d[lv�ri�v\pphwu\ fhphqw�sdvwh��� vwdwlrqdu\�phvk dlu���vwdwlrqdu\�phvk )j d� dlu�fhphqw�s��� lqwhuidfh dlu�fhphqw�s��� lqwhuidfh �vwdwlrqdu\�zdoo��qr�vols�e�f�� �vwdwlrqdu\�zdoo��qr�vols�e�f�� figure 1. schematic set up of mini-slump (symmetric part of the cone is shown) test with shown boundary conditions a) without mold b) with mold. in vof model volume fraction of phase in an individual cell is set to either 0 (none phase is present) or 1 (phase is fully present). in the sharp interface region, the value in the cell could range from 0 to 1. in order to preserve a sharp interface, one of few discretization schemes has to be used. the more details about discretization schemes could be found in ansys fluent theory guide [6]. in the current study, explicit scheme formulation with a geometric reconstruction scheme [8] was used to solve the volume fraction equation. 2.2. material model computationally, the concrete can be assumed as non-newtonian fluid. several models, such as bingham or herschel-bulkley, allow to model concrete flow. in both models it is required to overcome the yield stress threshold to initiate flow, the flow of fluid is further prescribed with plastic viscosity. for the bingham plastic model, after the stress is initiated, the shear stress is a linear function of the shear rate, and fluid behaves as newtonian. herchel-bulkley’s model exhibits shear-thinning behavior. in ansys fluent for shear rate γ̇ > critical shear rate γ̇c the viscosity is calculated as η = τ0 γ̇ + kγ̇n−1, (3) where τ0 is the yield stress threshold, k is the consistency index, and n is the power-law index. for bingham model n = 1, for herchel-bulkley: 0 < n < 1. critical shear rate γ̇c is the point beyond which material acts like very viscous fluid [6]. for γ̇ < γ̇c the viscosity is expressed by equation η = τ0(2− γ̇γ̇c ) γ̇c + kγ̇n−1c % (2 − n) + (n − 1) γ̇ γ̇c & . (4) 3. numerical modeling the presented numerical model was used to reproduce the slump flow test of mini-cone on the haegerman table described in en 1015-3. the dimensions of the mini-cone are defined as following: upper radius = 70 mm, bottom radius = 100 mm, height = 60 mm. the numerical model was solved as a 2d axisymmetric task. in general, two approaches could be followed to obtain slump spread and height of the mini-cone. in the first approach, the mold is not taken into account in the simulation. the mesh is stationary, and the overall calculation time is short. in the second approach, the mold and its lifting are added to the simulation, which requires dynamically changing mesh. the calculation time is then significantly prolonged. one set of parameters of the herschel-bulkley material model was used in the modeling to show differences in boundary conditions. the set of parameters was chosen for typical cement paste without admixtures as follow: τ0 = 25 pa, k = 10 pa·sn, n = 0.5, density ρ = 1800 kg/m3. critical shear stress was also fixed at a value γ̇c = 0.005 s−1. the air was modeled as newtonian fluid with density = 1.225 kg/m3 and viscosity = 1.7894e-05 kg/m·s. 3.1. slump test without mold the schematic situation of the model without mold is shown in fig. 1a. the domain of 300×100 mm was filled with linear quadrilateral cells. the frequency of cells near the base wall was increased to capture the velocity profile correctly, see fig. 2a. no-slip boundary condition was prescribed on the base wall. the initial trapezoidal shape of cement paste acts in the gravitational field, which is the only external force acting on the mini-slump. 3.2. slump test with mold lifting the scheme of the model with mold is shown in fig. 1b. the mold with thickness of 8 mm was added on the left side mini-cone, and its lifting was modeled 59 jiří němeček, jiří němeček acta polytechnica ctu proceedings ����pp ����pp ���pp ���pp ���pp ���pp ����pp ����pp ���pp d[lv�ri�v\pphwu\ �� �p p �� �� p p d[ lv �r i�v \p p hw u\ �� �p p �� �� p p &���vwdwlrqdu\�phvk�� %���vwdwlrqdu\�phvk�� $���g\qdplf�phvk�� d� e� � vwdwlrqdu\�phvk�� figure 2. mesh used for the slump test modeling a) without mold b) with mold. as rigid body motion in the vertical direction. several cases with various velocities: 0.01 m/s; 0.05 m/s; 0.2 m/s; 2.0 m/s were calculated to show differences in final slump spread. such a problem requires a dynamically changing mesh of the domain. triangular cells were chosen as most suitable for the analysis. in order to save calculation time, only one part of the domain was used for adaptive meshing. thus, the domain 300×200 mm was then divided into three regions. the first part (a) around the mold and cement domain was filled with the small size, 2.5 mm triangular cells. in this part, adaptive changes in the mesh: smoothing and remeshing were applied. the second part (b) is above the rest of the table with 2.5 mm cells . the cells were smoothened only above the table (wall b.c.). the third part (c) was filled with larger 15 mm cells. both the second and third parts of the mesh stayed stationary. the mesh is shown in fig. 2b. in the case of adaptive meshing in ansys fluent, the divergence problems occurred on the interface of mold and table (wall-wall boundary condition) where, after the initial lifting, the mesh is missing under the mold. to overcome such a difficulty 1 mm of mold was cut off in model, which should have minimal effect on the results. the gap was filled with triangular cells and air particles. the mold wallcement interface was modeled with two variants: a) no-slip b.c., b) specific shear b.c. with a value of 0.3 pa. 4. results and discussion fig. 3 shows three variants of the slump spreading results displayed in five different times. the first variant is the slump spread without mold (fig. 3a). after the calculation begins, all the volume of cement paste starts to flow. no obstacle prevents cement from movement, and due to the high inertia effect, the slump spread reaches the highest value (2×99.8 = 199.6 mm). the result with considering mold lifting with a velocity of 0.05 m/s and no-slip b.c. prescribed on the mold and table is shown in fig. 3b. obviously, the mold is partly holding the whole cement volume from motion, which is the factor decreasing the final spread (2×87.3 = 174.6 mm) which is 12.5% decrease in slump spread (199.6 mm) beside no-mold condition. due to no-slip b.c., part of the cement paste is sticked to the mold, which is another factor reducing the final spread. the third case (fig. 3c) shows mold lifting (v = 0.05 m/s) with shear b.c. prescribed on the mold, the chosen value 0.3 pa represent almost no shear between mold-cement interface. in this case, the cement is not sticking to the mold, which functions only as an obstacle. the overall result of slump spread (2×92.7 = 185.4 mm) is 60 vol. 30/2021 slump flow test numerical simulation ���pp w� ������v ���pp dlu fhphqw�s�� w� ��������vw� �����v ���pp ���pp w� ��������vw� �������v ���pp w� ��������vw� ������v w� ��������v ���pp w� ������v w� ������v dlu fhphqw�s�� ���pp d��qr�prog e��prog�oliwlqj� ������p�v��qr�vols�e�f� w� ������v dlu fhphqw�s�� ���pp f��prog�oliwlqj� ������p�v��vshfl¿f�vkhdu�e�f� w� ��������v ���pp ���pp w� ��������v �����pp ���pp �����pp ���pp ���pp ���pp w� ��������v ¿qdo�vsuhdg� �������pp ¿qdo�vsuhdg� �������pp w� ��������v ���pp �����pp ¿qdo�vsuhdg� �������pp figure 3. results of calculation of mini-slump test in five different times: a) without mold b) with mold lifting (v = 0.05 m/s) and no-slip b.c. c) with mold lifting (v = 0.05 m/s) and specific shear b.c. on the mold (0.3 pa). higher than in the previous case with no-slip b.c. but lower than in model without the mold. the decrease in the slump spread is only 7.1% besides the no-mold case. in terms of reproducing the real behavior of the mini-slump test, the case in fig. 3b is most promising. since during the experiment, cement paste is always sticking to the steel mold. applying the real value of friction would be even more suitable. however, establishing the coefficient’s value on a water wet steel-cement paste interface is not a simple task since the amount of water used, and admixtures used differs in each individual experiment. the results of the final spread and slump are summarized in tab. 1 and shown in fig. 4. it is apparent that lifting velocity affects overall results. basispread slump increase of [mm] [mm] spread v = 0.01 m/s 171.8 23.4 +0.0% v = 0.05 m/s 174.6 22.1 +1.6% v = 0.20 m/s 176.4 22.2 +2.7% v = 2.00 m/s – – – no-mold 199.6 17.2 +16.2% table 1. results of spread and slump calculated with different mold lifting velocities and no-slip b.c. cally, for lower lifting velocity (v = 0.01–0.20 m/s), the spread is increased when the lifting velocity is in61 jiří němeček, jiří němeček acta polytechnica ctu proceedings no mold v = 0.01 m/s v = 0.05 m/s v = 0.20 m/s 0255075100125 spread [mm] s lu m p [m m ] 10 20 30 40 50 0 figure 4. results of spread and slump calculated with different mold lifting velocities. creased. however, the difference between the smallest (v = 0.01 m/s) and maximum (v = 0.20 m/s) spread is only 2.7%. which is much lower than the difference 16.3% between mold and no-mold lifting. the high lifting velocities 2 m/s bring the cement paste to motion where other forces than gravitational are acting. the slumping process is not smooth, and the results are most probably incorrect due to the turbulent flow and calculation instability. similar findings were found in gao’s study, where different geometry and material were used [5]. the calculation time necessary for the model without a mold and stationary mesh is about 5 minutes. the time needed for calculation with dynamic mesh and mold lifting is 3-7 hours depending on the b.c. the total calculation time could be reduced using variable time step where the smaller time step is necessary at the beginning of the calculation and could be further increased during the calculation. 5. conclusions the present work provides a general study of setting up a numerical model of slump flow test and shows how different boundary conditions could affect the results. the calculation was done on typical cement paste using only one set of parameters of the herchelbulkley model to neglect different mixtures’ effect. several conclusions are made. • the consideration of the mold lifting during calculation has a significant effect on results. the final slump spread is increased by 12–16 % when the mold lifting is not used in the calculation. • mold lifting velocity has only a small effect on overall slump spread while slow lifting velocities are used. slow lifting velocities are considered in the range of 0.01–0.20 m/s, which should be typical lifting velocities used in laboratory conditions. the difference in slump spread between minimum and maximum velocity was calculated as 2.7 %. • lifting velocities with higher than 2 m/s bring fluid to turbulent motion, which is not recommended to calculate in ansys fluent with the herchelbulkley model suitable for laminar flow. moreover, these velocities are not commonly used in the experiments. • consideration of mold lifting and dynamic remeshing in the calculation results in increasing calculation time from minutes to hours. acknowledgements financial support of the technological agency of the czech republic (trend fw-01010521) and the grant agency of the czech technical university in prague (sgs20/107/ohk1/2t/11) are gratefully acknowledged. references [1] c. ferraris, f. de larrard, n. martys. fresh concrete rheology: recent developments. materials science of concrete vi, amer cer soc ed s mindess, j skalny pp. 215–241, 2001. 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[8] d. youngs. time-dependent multi-material flow with large fluid distortion, vol. 24, pp. 273–285. 1982. 62 https://doi.org/https://doi.org/10.1016/j.cemconres.2010.06.005 https://onlinelibrary.wiley.com/doi/pdf/10.1002/nag.419 https://doi.org/10.1002/nag.419 https://doi.org/10.1122/1.1879041 https://doi.org/https://doi.org/10.1016/j.mineng.2014.11.001 https://doi.org/https://doi.org/10.1016/b978-0-12-815485-4.00018-8 https://doi.org/10.14311/app.2022.33.0316 acta polytechnica ctu proceedings 33:316–321, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague railway reinforced concrete infrastructure life management and sustainability index seyed mohammad sadegh lajevardi∗, paulo b. lourenço, hélder s. sousa, josé c. matos university of minho, isise, department of civil engineering, campus of azurém, 4800-058, guimarães, portugal ∗ corresponding author: sadegh.lajevardi@gmail.com abstract. infrastructure healthy enhancement for saving resources in operation procedures is one of the most important objectives for owners on their decision support system based on cost management. in this manner, finding the intervention action priority, as well as the inspection method and maintenance system for each component, with regard to a limited resources amount is investigated in this paper. defects on infrastructure components create data and these data are undoubtedly useful to increase the knowledge in decision making in practice. in that sense, risk analysis and value of information can be applied using decision trees together with bayesian networks. for data filtering and noise reduction, a principal component analysis may also be applied to manage a database and prepare useful input variables for the decision tree system. this paper presented an approach for the maintenance managers to prepare their infrastructure available with a sustainable index with minimum cost. this index would be a tool for decision-makers with regard to the cost management and sustainability aspects. keywords: life management, maintenance planing, sustainability index. 1. introduction most of the transportation modes are made through railway networks and road infrastructures. passengers’ expectation is directly affected by the quality of structures and equipment during the operation. saving quality level requires planning and predicting resource consumption based on the status of their materials in their degradation model and maintenance approach. the maintenance approach would be renewed or retrofitted during the operation. quality will explain by performance indicator which has been affected by several attitudes, especially peripheral climate such as strong corrosion processes in the coastal area or freezes and thaw in deserts that result from temperature degree changes or moisture percentage changes in terms of sustainability and environments [1]. therefore, there is a wide range of alternatives and conflicting criteria are involved as the root of failure on infrastructure. therefore, multi-criteria decision making (mcdm) has been applied to estimate the infrastructure’s element performance during the operation [2]. recent research, for finding the sustainability index has been focused on several subsets of mcdm to prepare a decision support system such as preference selection index (psi) [2, 3]. this method is typically chosen when there are no significant weights for attributes. in other research, in order to mine the effective parameters for the indexing model based on pareto concept and making pair-wise comparison questionnaire serving analytical hierarchy process (ahp) requirements to find the weights of them [4]. ahp method has been adopted for the calculation of the sustainability assessment index (sai) based on normalized key performance indicators and their weight, which has been calculated by ahp [5]. the theoretical framework developed for extracting the sustainability index (si) is a function that comprises wellbeing, resource, compliance, and resources [6]. for qualitative assessment in the ahp process, recent research has been used the likert scale and extract sustainability level [7]. the results led to the development of a sustainability index model (sim) useful to assess manufacturing performances or maintenance procedures during the operation with rank alternatives [8]. also, these results have been developed with respect to the evaluation criteria and the weights of the criteria to prioritize the network systems [9]. these indexes could be extracted for prediction by neural network method to decision making with considering probable scenarios [10]. since, future generations are important for these studies, prediction, and monitoring of the results with any scenario is necessary for decision-makers. therefore, sustainable development is a comprehensive solution for the present and future of humans [11]. based on recent research, the sustainable index is an important index for the bridge management system and green infrastructures and green cities. 2. aim of the research this paper presents an approach to developing a practical indexing model in terms of resource management 316 https://doi.org/10.14311/app.2022.33.0316 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 railway reinforced concrete infrastructure subject extracting indexmethod subset method coastal climate adaptation planning and evolutionary governance: insights from homer, alaska. [1] interviews with key informants development of sustainable performance index (spi) for self-compacting concretes (scc) [2] mcdm spi application of preference selection index method for decision making over the design stage of production system life cycle [3] mcdm psi sustainability index for highway construction projects [4] mcdm ahp development of sustainability assessment index for machine tools [5] mcdm ahp development of a multidisciplinary approach to compute sustainability index for manufacturing plants singapore assessing the feasibility of using the heat demand-outdoor perspective temperature function for a long-term district heat demand forecast [6] mcdm ahp groundwater sustainability assessment framework: a demonstration of environmental sustainability index for hanoi [7] mcdm ahp developing a sustainability index for mauritian manufacturing companies [8] rank correlationordinal association kendall coefficient life cycle aggregated sustainability index for the prioritization of industrial systems under data uncertainties [9] lcc net present value (npv), internal return rate (irr), and payback time (pt)) . predicting subjective measures of walkability index from objective measures using artificial neural networks [10] ann comparison of sustainability models in development of electric vehicles in tehran using fuzzy topsis method [11] mcdm fuzzy topsis the lisbon ranking for smart sustainable cities in europe [12] featureselection principal component analysissensitivity analysis table 1. research method and index extract. and optimization. it will prepare a decision support system for managers that leads to wider uptake welfare during the infrastructure life cycle for their real owner, people. since this plan optimizes resource consumption, therefore it will keep the environment safe and prevent nature from contamination. based on this framework, the manager will make an optimized decision for the material and shape of their structures’ element during the design. this framework would also be useful for maintenance managers to decide about their maintenance plan with a comparison between the elements and components of the railway network. 3. research contribution this paper is structured as follows: • data collection from overall monitoring, such as visual inspection. • variance measurement and extracting the data value • defining the reliability index • define the effective parameters in terms of sustainability with a questionnaire form • correlation analysis between the reliability index and sustainable parameters • prepare a sustainable model for maintenance managers 317 s. m. s. lajevardi, p. b. lourenço, h. s. sousa, j. c. matos acta polytechnica ctu proceedings ��������� �� ��� �������� �� �� �� ��� � �� ��� �������� ��� � �� ��� �� ����� � �� �� ������ ��� �� � � ���� ������ � ��� ��� ���� � ��� �� ���� � ��� ��� ������ � �� ��� �� ����� ���� ��� ���� ������ ����� ���� ��� ��� !� � �� ��� ������� �� !�!��� � �� ��� �� �� � �� �� �� ��� ��� ����!�� ����� �� ������ ���� ������ �������� "��� �#��� figure 1. flow diagram of extracting sustainable index. 4. methodology this research would establish a comprehensive mathematical method for extracting the sustainability index for infrastructure as follows. • defect detection and extracting the probability of failure • principal component analysis (pca) and the value of information for reducing the useless data • environment features and extracting the sustainability parameters • replacing the roots of the defect (environment features) by the defect and extracting the sustainability index for case study • the weight of each parameter shows the importance of the index elements. based on these weights, a new construction plan and materials regarding their maintenance method will be updated. the framework of extracting sustainable index has been shown as follows diagram. 5. data collection this research would be establishing a comprehensive mathematical method for extracting the sustainability index to assess infrastructure in terms of sustainability. this research has been focused on the tehran subway bridge for monitoring and quality assessment during the operation. based on the inspection checklist, five bridges are determined and inspected as a sample by expert inspectors. these bridges have been assessed by the items of the national code [14]. parameters with subjective attitudes have been ranked based on their environmental status and the other geotechnical features between 1 to 4 or 1 to 5. some of these parameters would be determined according to the observed defects in each element of the bridge. since these parameters are not the same range, they will be normalized by formula as follows. z = x − µ σ (1) in this formula, µ is the arithmetic mean and σ is the standard deviation of the distribution. row data after normalization process transfer to the principal component analysis (pca) framework. based on the scree plot and their eigenvalue, three components have the most value of information. observation from inspection determines the status of the bridge’s elements as follows. some of these elements related to defects and they were affected by environments’ features according to the last table. sustainable structures during their life cycle have to fulfill the thresholds regarding their components and the environment’s features based on their weight. by considering these tools, it will be possible to manage assets along their lifetime in a more sustainable and efficient way [15]. therefore, pca has been used 318 vol. 33/2022 railway reinforced concrete infrastructure component 1 2 3 liquefaction (li) .825 .448 -.084 land slide (ls) -.012 -.681 .258 rock fall (rf) -.811 -.252 .031 thunderstorm (th) .313 .034 .931 flood (fl) .675 -.475 .484 climate (cl) .825 .448 -.084 density of soil (ds) -.558 .648 .306 underground void distance (ud) -.553 .506 .508 table 2. component matrix. defects type dependent element cl fl th ls li ds ud rf deck stability deck • • drainage drainage/pier/deck/ foundation • • • foundation/embankment settlements foundation/abutment • • • • • • railway track geometry defect (gage widening, alignment, twist, longitudinal level [13]) railway track • bridge thermal movement deck/railway track • temperature stress cycling all bridge’s element • deterioration all bridge’s element • scour pier/foundation • • table 3. impact of environment features on bridge. bridge number defect onrailway track defect on deck defect on abutment defect on pier defect on foundation 1 6 0 0 0 0 2 3 1 0 0 0 3 7 4 1 0 0 4 1 3 0 0 0 5 0 4 1 0 0 6 1 3 0 0 0 7 1 0 0 0 0 table 4. number of registered defects based on bridge’s elements. bridge number number of defect on element [table 4] defect type approximate volume (m3) probability of defect [table 4] (%) impact [table 2 & 3] sustainability index 4 3 instabilityof deck 709.8 0.42 0.931rf+.031th 0.42(0.931 rf+.031th) table 5. sustainability index for a bridge in case study. 319 s. m. s. lajevardi, p. b. lourenço, h. s. sousa, j. c. matos acta polytechnica ctu proceedings figure 2. scree plot for component’s eigenvalue based on pca analysis. and 3 components have been extracted by the pca method for 8 environment features with their weights as follows. based on table 2, li represents the probability of liquefaction and also other environmental features such as ls and rf and the other. this method will keep the valuable data and reduce the main by ranking the parameters in each component [16]. it is necessary for monitoring the mean value and variance changes for each feature, which would be important for sustainable structures. for example, increasing the mean value for flood shows the increasing probability of the scour surrounding the piers of the bridge. since climate change would be effective in increasing the crack growth of concrete elements, these features would be related to these types of defects. these relations are shown as follows. for finding the sustainability index, it is necessary to find the risk according to weight for their defects. the probability of failure will be extracted as the following formula [17]. for each bridge (see table 4), v is an approximate measure representative of their volume regarding length (l) and width (w). the number of defects in each element has been determined by visual inspection [18] (f) based on table 4. this shows that a larger bridge has more capacity for a defect in its components when compared with a smaller one. probability of failure (p-f) are normalized by dividing them with the average of all failure densities. pf = ! f" v dv (2) v ≂ l × v (3) based on table 3 and table 2, the sustainability index has been presented as follows. bridge number 4 in this step is a sample for calculating the sustainability index for the instability of the deck. with this result, it is necessary to consider the risk of reinforced concrete elements and their defects during the operation with regard to rockfall and thunder. it is obvious for other bridges in the same area, impact factor is the same but the probability of defect will change based on their structural features. therefore, the sustainability index will change for each bridge, and the final grade will determine by their environment features and structures. 6. conclusion the sustainability index would be a combination of environmental features to make a decision during the operation or designing of the structure. in this approach after putting up the probability of environment features such as rockfall or climate, finding the location, material and bridge structure type would be possible by benchmarking and expert judgments. with this index it is possible to locate the bridge with a minimum life cycle cost and maximum sustainability. finding the optimum plan with regard to climate change and seasons is also possible during the operation. for example, in winter railway track face with a wide range of changes in some area and this issue consequently effects on increasing the railway track failure such as rail and fastening system breakage. therefore, finding the proper place, the best material and optimum maintenance plan would be possible with the sustainability index. acknowledgements acknowledgments this work was partly financed by fct / mctes through national funds (piddac) under the r&d unit institute for sustainability and innovation in structural engineering (isise), under reference uidb / 04029/2020. this project has received funding from the european union’s horizon 2020 research and innovation programme under grant agreement no769255. the sole responsibility for the content of this publication lies with the authors. it does not necessarily reflect the opinion of the european union. neither the innovation and networks executive agency (inea) nor the european commission are responsible for any use that may be made of the information contained therein. references [1] s. 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orientations david vokoun∗, jan maňák, karel tesař, stanislav habr czech academy of science, institute of physics, na slovance 1999/2, prague, czech republic ∗ corresponding author: vokoun@fzu.cz abstract. the thermomechanical processing by equal-channel angular pressing (ecap) is used for certain metals and alloys in order to make their structure fine and to increase material strength. in the previous study done at our institute, grade 2 titanium was successfully processed using four consecutive route a passes via a 90 ◦ ecap die with high backpressure at room temperature. orientation dependence of compressive and tensile loading of ecaped titanium samples was demonstrated at macro-scale. however, scarce attention has been paid so far to the mechanical behavior of ecaped titanium samples at micro-scale. in the present study, compression experiments on titanium micropillars, fabricated using focused ion beam, are carried out for two main directions in respect to preceding ecap pressing (insert and extrusion directions). the purpose of this study is to discuss the orientation dependence of mechanical response during compression of the as-ecaped titanium micro-pillars. keywords: ecap, micro-pillar, titanium. 1. introduction titanium is a light-weight material with hexagonal structure at room temperature. due to high specific strength (strength-to-weight ratio), titanium and titanium based alloys are used in carand aerospaceindustries. in addition, titanium and its oxides are biocompatible. sound biocompatibility together with high resistance to the onset of localized corrosion make titanium a good candidate for biomedical applications. furthermore, fabrication of metalic micro-parts has been receiving much attention in recent years due to the rapid development of micro-electro-mechanical systems (mems). therefore, understanding of mechanical properties of various metalic materials including titanium at micrometric level is of importance. it only makes sense to produce and mechanically test microparts which are either monocrystalline with a fixed relation between the crystallographic and geometric axes [1] or which contain submicron grains (grains which are small enough to guarantee reproducibility of mechanical tests carried out on arbitrary two microsamples of identical dimensions). however, grains of titanium samples produced in a conventional way are large and their refinement requires special methods. the small tensile strength of titanium (commercially pure, grade 2 titanium, has tensile strength only 345 mpa [2]) can be increased either through alloying [3, 4] (a high chance of decreasing or losing biocompatibility), or through severe plastic deformation (spd) processes used for grain refinement. the increase of tensile strength in various metals and alloys through ecap (equal channel angular pressing) and other spd methods is generally known, e.g. the authors of [5] showed an increase of tensile strength of ecaped commercially pure (cp) titanium due to the ecap process up to 1000 mpa (true stress). ecap is a wide-spread spd technique [6] used also for titanium and titanium based alloys. during an ecap procedure, samples (billets) are heavily deformed with strains close to or exceeding 100% (depending on the used die geometry and the number of passes) due to repeated pressing a metalic billet through a system of two channels, axes of which intersect under an abrupt angle, usually 120◦ or 90◦. a great number of studies, so far published, describe preparation of ecaped titanium samples and evaluate microstructure, various physical and mainly mechanical properties of ecaped titanium samples in dependence on the ecap parameters [5, 7–11]. the key ecap parameters are related first to the orientation of a billet upon the repeated insertion in a die in respect to the previous passing of the billet through the die (routes a, ba, bc, and c [6]). the ecap parameters are also related to the die geometry and the billet temperature. angle ϕ (channel angle) is the abrupt angle between the two die channels and angle ψ (corner angle) represents the outer arc of curvature where the two channels intersect [6]. the other ecap parameters are e.g. number of the passes through a die, pressing speed, values of backpressure, pressing temperature, etc. room-temperature ecap processing is not so common for titanium although it brings the advantage of enhanced grain refinement due to the suppressed grain structure recovery. room-temperature ecap processing using channel angle 90◦ and less with cp grade 2 titanium is still challenging [5] due to material cracking after the first or second pass of the material through the die. in ref. [5], room-temperature ecap processing is described and later used in our study in exactly the same way with the same titanium alloy (identical chemical composition). 1 http://dx.doi.org/10.14311/app.2020.27.0001 http://ojs.cvut.cz/ojs/index.php/app d. vokoun, j. maňák , k. tesař, s. habr acta polytechnica figure 1. scheme of the ecap die with the coordinate system. the abbreviations id, ed and td represent the insert, extrusion and transverse directions, respectively. figure 2. the sem image of a micro-pillar (a) and the diamond tip approaching one of the fabricated micro-pillars (b). the purpose of the present work is to extend the previous work [5] in relation to compression experiments at room temperature. to our knowledge, so far no work on compression of ecaped titanium micropillars has been done. however, mechanical testing on micro-pillars of other ecaped metals and alloys has been performed recently [12, 13]. the advantage of our work is the availability of compressive tests on ecaped titanium at both macroand micro-scale and the possibility of comparison. 2. samples a cube, 3 mm × 3 mm × 3 mm, made of ecaped titanium was cut out (using wire electro-discharge machining) in such a way that each cube face was perpendicular to one of the principal directions insert direction (id), extrusion direction (ed) and transverse direction (td) (see fig. 1). the micro-pillars, each in the shape of a rectangular prism with a square base (see fig. 2a), were micro-fabricated and observed using an fei quanta 3d dual-beam sem/fib (scanning electron microscope/focused ion beam) system (fei, hillsboro, usa) with a ga+ ion source operated at 30 kev with various currents using automated milling script. each micro-pillars’ face was again perpendicular to one of the principal directions (id, ed, and td). as for dimensions of the micro-pillars, their height and width were approximately 8.5 µm and 3.5 µm, respectively. the micro-pillars were of two types denoted as id and ed micro-pillars. id {ed} micro-pillars had their long axis parallel to id {ed} direction, respectively. no micro-samples with pillar’s long axis parallel to td direction were prepared since previous studies showed almost identical mechanical characteristics in id and td directions. it was checked that each micro-pillar contained grains small enough not to limit reproducibility of the following compression tests. there were three id and four ed micro-pillars fabricated; together, seven micro-pillars. for each micropillar just one force-displacement curve was measured and converted into stress-strain curve. however not all 2 vol. 27/2020 ecaped titanium micro-pillars compression stress-strain data were used. most stress-strain curves were excluded because of a misalignment during the compression. 3. methods the ecap die with a channels’ diameter of 11 mm and the channel and corner angles equal to 90◦ and 0◦, respectively, was used. back pressure regulated between 270 mpa and 590 mpa to maintain an average plunger pressure of 2800 mpa was applied during the ecap process. the inserted titanium rod-like sample, 35 mm long, was encapsulated in a lubricant-treated thin steel container in order to eliminate frictional forces. the details of the ecap process can be found in [5]. the ecap process was carried out at room temperature using route a. the number of passes of the billet through the die was four. the in situ compression tests were performed at room temperature using a hysitron pi 85 sem picoindenter (hysitron, minneapolis, usa) with a 20 µm flat-punch diamond tip. figure 2b shows the diamond tip approaching one of the fabricated micro-pillars. the true strain rate was around 5.2 × 10−4 s−1. the micro-pillars were compressed up to about 10 %. a special care was given to the alignment of the loading and micro-pillars’ long axes during loading. the loading data with an obvious difference in young’s moduli at the onset of loading and unloading were excluded. the grain morphology at the micro-pillar longitudinal sections were acquired using the fei quanta 3d equipped with an edax hikari ebsd (electron backscatter diffraction) detector (edax, new jersey, usa) and analyzed with a help of edax oim analysis v8 software. 4. results and discussion the grain morphology of the micro-pillars (not shown here) corresponds to the micro-structure of the ecaped titanium samples studied in [5]. herein, it is worth mentioning that the mean grain size determined in [5] using an x-ray diffraction (xrd) (crystallite size calculated from line broadening of the xrd peaks) was 50 nm and the prevailing grain size determined in [5] using a transmission electron microscope was in the range 50 nm – 300 nm. the ebsd map of an ecaped ti sample processed identically as the present sample can be found in [14]. as of the present micro-samples, there are regions of recrystallized grains with a mean grain size of 200 nm and regions of non-recrystallized grains. the grain size of the largest non-recrystallized grains does not exceed 3 µm. the micro-samples contain only a few large grains and these grains still contain regions with various crystallographic orientations. in addition, an arbitrary micro-pillar cross-section area (3.5 µm × 3.5 µm) contains at least five grains. low-angle grain boundaries prevail among the noncrystallized grains. furthermore, it can be assumed that high deformation energy is accumulated in the regions of non-recrystallized grains. figure 3 shows the diagram of true compression stress versus true strain for id and ed micro-pillars, id and ed macrosamples and titanium bulk sample. in fig. 3, the stress-strain data belonging to compression loading of ecaped titanium samples denoted as id and ed macro-samples and titanium macro-sample with no ecap treatment denoted as ti bulk (coarse-grained sample) were taken from our previous study [5]. the id, ed and ti-bulk macro-samples had a shape of a cylinder with a diameter of 3 mm, a height of 6 mm and their rotational axes along id, ed and ld directions, respectively. all the samples were mechanically tested at room temperature. in [5], it is shown that at macroscopic level, mechanical properties of ecaped titanium are anisotropic. the maximum compression stress in id direction is about 24 % higher than the maximum compression stress in ed direction (see fig. 3). figure 3. the diagram true compression stress versus true strain for id and ed micro-pillars, id and ed macro-samples and titanium bulk sample. at microscopic level, in case of id and ed micropillars, maximum stress in id micro-pillar is still greater than maximum stress in ed micro-pillar but the stress ratio, (σid)max/(σed)max, is close to 1. there is a question whether the micro-samples are less anisotropic (in compression) than the macro-samples. however, at this stage of our study, it is not possible to make any conclusion. stress-strain data for more micro-pillars are necessary in further research. if it turns out that there really is less compression anisotropy in the micro-pillars compared to the macrosamples, then a potential explanation might be as follows: the effect of anisotropy in compression, i.e. yield strength depends on the compression direction, is likely due to anisotropic distribution of dislocations in the ecaped samples. if the size of the ti ecaped samples decreases, as in the present study, to units of microns, then the plastic deformation (during compression) is associated with the creation and movement of newly formed dislocations, rather than with motion and interactions of already existing dislocations (as it is in bulk samples) [15]. this means there is a weaker 3 d. vokoun, j. maňák , k. tesař, s. habr acta polytechnica effect of the existing anisotropic distribution of dislocations. this might be a reason for the potentially less significant effect of compression anisotropy in the micro-pillars, compared to the macro-samples. another distinct phenomenon in fig. 3 is the samplesize effect, obvious when comparing the true stress true strain curves belonging to the micro-pillars and the ecaped titanium macro-samples. it is assumed that the mean grain size of the compared samples is approximately identical. xu et al. [16] observed the sample-size effect for pure aluminum in the form of softening behavior (negative work-hardening rates) and a non-significant yield stress change with a reduction of the specimen size. however, for the fixed small grain sizes, the authors did not test sample dimensions for which the sample-size effect is significant i.e. when the sample size decreases to dimensions comparable to about ten times the average grain size. in addition, the crystal structure of aluminum is cubic whereas the crystal structure of titanium is hexagonal at room temperature. in the present study, both yield stress and the work-hardening rate increased remarkably with the reduction of sample dimensions to only units of microns. these observations are perfectly consistent with compression tests done on ecaped mg-5 wt.% al micro-fabricated into a shape of micro-pillars [17]. the sample-size effect observed in the present study is based on the dislocation starvation theory [15] and the effect of grains close to the sample’s surface since the volume ratio of surface grains increases with the sample size decrease. the theoretical background of the observed sample-size effect might be found also in [18]. 5. conclusions ecaped titanium (grade 2) is successfully reproduced at room temperature (route a, four passes) using an ecap-back-pressure die with a channels’ diameter of 11 mm and the channel and corner angles equal to 90◦ and 0◦, respectively. a concern was raised that ecaped titanium micro-samples in compression might be less anisotropic than the macrosamples. further study is necessary to clarify this concern. furthermore, comparing the compression stress – strain curves for ecaped titanium microsamples and macro-samples, both yield stress and the work-hardening rate increase remarkably with the reduction of sample dimensions. acknowledgements support of the work under the project ga17-05360s is acknowledged. references [1] j. maňák, d. vokoun. microbending experiments on pure magnesium with nonbasal slip orientation. materials 11(8):1434, 2018. doi:10.3390/ma11081434. 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[14] h. sopha, k. tesar, p. knotek, et al. tio2 nanotubes grown on ti substrates with different microstructure. materials research bulletin 103:197–204, 2018. doi:10.1016/j.materresbull.2018.03.036. 4 https://doi.org/10.3390/ma11081434 https://doi.org/10.1016/s0921-5093(97)00806-x https://doi.org/10.1016/s0924-0136(96)00030-1 https://doi.org/10.2320/matertrans1989.34.1217 https://doi.org/10.1016/j.msea.2015.07.038 https://doi.org/10.1016/j.pmatsci.2006.02.003 https://doi.org/10.1016/j.jmatprotec.2014.12.024 https://doi.org/10.1016/j.msea.2014.04.014 https://doi.org/10.1016/s1359-6454(02)00501-3 https://doi.org/10.1016/j.msea.2011.10.062 https://doi.org/10.1080/10667857.2016.1238131 https://doi.org/10.1016/j.msea.2015.09.110 https://doi.org/10.1016/j.actamat.2019.08.059 https://doi.org/10.1016/j.materresbull.2018.03.036 vol. 27/2020 ecaped titanium micro-pillars compression [15] c. frick, b. clark, s. orso, et al. size effect on strength and strain hardening of small-scale [111] nickel compression pillars. materials science and engineering: a 489(1-2):319–329, 2008. doi:10.1016/j.msea.2007.12.038. 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[18] v. kafka, d. vokoun. nanowires: diameter-dependent strength criterion. iop conference series: materials science and engineering 504:012110, 2019. doi:10.1088/1757-899x/504/1/012110. 5 https://doi.org/10.1016/j.msea.2007.12.038 https://doi.org/10.1016/j.msea.2015.08.060 https://doi.org/10.1109/inec.2010.5424788 https://doi.org/10.1088/1757-899x/504/1/012110 acta polytechnica 27(0):1–5, 2020 1 introduction 2 samples 3 methods 4 results and discussion 5 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0020 acta polytechnica ctu proceedings 33:20–26, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague using an analysis of concrete and cement epd: verification, selection, assessment, benchmarking and target setting jane anderson∗, alice moncaster open university, walton hall, milton keynes, keynes mk7 6aa, uk ∗ corresponding author: jane.anderson@open.ac.uk abstract. the carbon embodied in buildings is an important proportion of our emissions and needs to be radically reduced in order to support climate change mitigation. the highest proportion of embodied carbon is usually emitted during the product stage, and within the structural elements. therefore, reducing the carbon embodied in the structural materials is likely to have a major impact. in most buildings, the majority of embodied carbon comes from steel and concrete. but although there are now hundreds of registered environmental product declarations (epd) for cements and concretes, there has been very limited independent published information comparing the embodied carbon of different concrete mixes and raw materials. this lack of comparative data limits the potential to make appropriate decisions at early design stages leading to low carbon buildings. the authors have recently conducted a review of verified epd for concrete mixes and for concrete’s key constituents, including cement, identifying the range of carbon coefficients. this paper provides guidance on making use of the coefficient ranges provided in that research: to support the verification of epd for concrete and its raw materials; in material selection; in assessing building level embodied carbon; in benchmarking; and in the setting of reduction targets. keywords: benchmarking, embodied carbon, epd, lca, methodology. 1. introduction and theoretical background the carbon embodied in buildings, from the use of construction materials which produce carbon dioxide and other greenhouse gases during their life cycle, is an important proportion of our carbon emissions. international energy agency (iea) estimate that embodied carbon accounted for 11% of global co2 emissions in 2017, most of this from cement and steel manufacture [1]. iea and the world business council for sustainable development cement sustainability inititive (wbcsd-csi) estimate the greenhouse gases emissions from cement manufacture are already 7% of our total global emissions [2]. iea also note that cement demand has nearly doubled between 2000 and 2016, mainly due to increased demand in china, and latterly, india [1]. with almost all cement used in concrete, it is clear that concrete has a major influence on global greenhouse gas emissions. environmental product declarations (epd) are standardised presentations of the environmental impact of products. iso 21930:2007 [3] was widely used by many of the epd programmes across europe, but en 15804 [4] has been developed following a mandate from the european commission to harmonise the provision of epd for construction products across europe, ensuring that epd to en 15804 have a common methodological approach to lca and provision of environmental impacts. iso 21930 was updated in 2017 to closely align with en 15804 [5]. constructionlca’s infographic [6] shows that the numbers of epd using en 15804 have risen rapidly with over 7000 epd available globally at the start of 2020. as a significant construction material, epd have been produced for both cement and concrete products, including both ready-mix concrete and precast concrete. as detailed in their paper [7], anderson and moncaster made a systematic review and analysis of all the published epd globally for ready-mix concrete and its constituent materials (cement, aggregates and admixtures) at the end of 2019. they found 252 epd for ready-mix concrete reporting data for over 2000 individual products, 108 epd for cement and cementitious materials covering 118 products, 88 epd for aggregates covering 117 products and 9 epd for specialist aggregates, and 16 epd for admixtures coming from 25 epd programmes in total. these epd all represent products available in the market including both national and regional industry average data provided by trade associations and manufacturer specific data for both average and specific products, and many of these epd, especially those for ready-mix concrete, provided individual data for more than one product, in some cases providing data for over 100 individual products in one epd. these epd were reviewed in terms of the data on "cradle to gate" embodied carbon coefficients (the global warming potential impact indicator for mod20 https://doi.org/10.14311/app.2022.33.0020 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 analysis of concrete and cement epd ules a1-a3 within epd) and their embodied energy (the primary energy and secondary energy resource indicators for modules a1-a3 within epd), together with other data such as clinker and cementitious content for cements and 28-day strength for concretes, where they are provided. this paper describes how the information reported in the systematic review and analysis can be used in epd verification, materials selection, embodied carbon assessment at building level, in benchmarking and in setting improvement targets. 2. previous studies hammond and jones provided an overview of the embodied energy and carbon coefficients reported in academic and industry literature in their inventory of carbon and energy database (ice database) [8]. they found 112 datapoints for ready-mix concrete, 92 datapoints for cement and 36 datapoints for aggregate, though in relation to embodied carbon they note that "there is often an absence of such data". damineli et al. reviewed 156 randomly selected papers from 1988-2009, 59 from brazil (covering 604 concretes) and 97 international (covering 981 concretes), considering the embodied carbon of different concretes together with binder content, and 28 day compressive strength [9], and this work was updated by scrivener et al. including examples of recent developments, including data from concretes formulated with up to 70% replacement of binder by filler [10]. the ice database was updated in 2011 [11] using only 3 new datapoints for cement, aggregates and ready-mix concrete. the ice database was updated again in 2019 [12] using embodied carbon data from published epd. this version used 22 datapoints for ready-mix concrete (16 ready-mix sources listed), 14 datapoints for cement (4 sources listed) and 164 datapoints for aggregate (22 sources listed). pomponi and moncaster reviewed carbon coefficients for cement and concrete in academic literature, identifying 58 coefficients across the two products [13]. van den heede and de belie similarly identified 12 datapoints for cement [14], salas et al., reviewed the literature on lca for cement and identified 16 embodied carbon datapoints [15], ganassali et al. identified 32 epd for cement [16], kurda, silvestre and de brito identified 17 cement datapoints and 20 aggregate datapoints [17] and braga, silvestre and de brito identified 16 concrete datapoints [18]. passer et al. have written about how epd are being used in the european market [19] and jelse and peerens about how they can be used for green public procurement [20]. ganassali et al. discuss how they produced benchmark values for cement using 32 epd, using the median value of the range for the reference value, with limit and target values set using the boundaries of the upper and lower quartiles of the range distribution [16]. jones has developed embodied carbon benchmarks using the arithmetic mean of datapoints from selected epd and others sources for each product [12]. 3. description of the systematic review and analysis of epd and summary of its results 3.1. methodology the review was conducted at the end of 2019. during the data collection phase, epd were downloaded from all the known epd programmes, using ecoplatform [21] and the north american pcr catalogue [22] and a general literature search. in the categorisation phase, epd were categorised into cement, different concrete mixes, aggregates, and dividing the epd which report multiples into a set of individual results. in the third phase, the data was extrapolated from the epd cataloguing and tabulating like with like. the data included the embodied carbon coefficients (gwp for a1-a3) together with other relevant data for all the separate products contained within each product group. in the fourth phase, visualisations were developed and in the fifth phase, the results were checked and verified. 3.2. cement embodied carbon coefficients figure 1, taken from anderson and moncaster [7] which explains some of the reasons for variation, shows the embodied carbon coefficients for all the identified cement epd plotted on a graph against clinker content. the types of cement (e.g., cem i, cem ii etc) are differentiated using colour. 3.3. ready-mix concrete embodied carbon coefficients figure 2 from the same source [7], shows the ranges for embodied carbon coefficients for all the identified products within ready-mix concrete epd plotted against the 28-day strength where it has been provided. epd ranges are identified based on the country of concrete production, with average epd produced by national trade associations highlighted as a shaded range, labelled, e.g., "french generic". this clearly shows that the ranges of impact vary significantly between countries, with french and german industry generic epd having much lower impacts for the same 28-day strength than the corresponding us and canadian; and all epd from mexico and saudi arabia having much higher impacts than those epd from australia and the u.a.e. for example. 4. using the embodied carbon coefficient ranges for cement and concrete in the process of verifying the results, the authors found in several cases datasets which were significant 21 jane anderson, alice moncaster acta polytechnica ctu proceedings outliers, and in these instances contacted the epd programmes to check if there were possible errors in the data provided. in these cases, several errors were identified and epd have been reissued. this led the authors to consider how the data collected in the systematic review and analysis could be used. 4.1. in epd verification the above example clearly illustrates how the systematic review and analysis can be used to check epd for cement, concrete and aggregates to identify if the results are reasonable. for cement epd, figure 1 can be used to check the clinker content and embodied carbon are within the expected region of the graph, especially for cem i, ii, iii and iv cements. the systematic review analysis also provided a box and whisker graph showing the range of embodied carbon for epd by country of production and by type of cement (cem i, cem ii etc), not included here, allowing verifiers to check the plausibility of embodied carbon results. additional graphs are provided in the systematic review [7], but not included here, for cem i and cem ii cement epd showing the embodied carbon broken down by co2 from calcination and from fuel use, and the primary and secondary fuels use, broken down by renewable and non-renewable sources, which again can be used to consider the plausibility of data provided in the epd for verification. for concrete epd for verification, we recommend checking them against figure 3 which provides the embodied carbon ranges (mean and upper and lower quartile) for different countries and 28-day strengths to see that the embodied carbon is in the expected region of the graph. this graph also shows how many epd results were included in each grouping. 4.2. in material selection cement selection must be considered alongside the functionality of the concrete that is required. figure 3 shows that that there is great variation in impact reported for any given strength, but that generally, there is an increase in embodied carbon per m3 as the 28-day strength of the concrete is increased. however purnell cautions against the selection of low strength concretes over high strength concretes on the basis of their reduced embodied carbon, highlighting that significantly more of a low strength concrete may be required to fulfil a particular function [23]. a similar conclusion was drawn by damineli et al. who showed that assessing co2 per m3 and per unit of structural performance (kgco2e/m3/mpa) suggested that c50 concrete was the optimal choice ide greater savings in embodied carbon (20-35%) than those achieved by replacing cement with pulverised fuel ash (10-25%) for example [23]. carbon intensities for ready-mix concretes with 28-day compressive strength over 10 mpa (from [7]) are shown in figure 4. many manufacturers are now able to provide epd for the range of concretes that they product, for example using national tools such as the french betie epd tool [24] or industry association tools such as the csi epd tool developed for wbcsd csi [25] or that developed for the norwegian ready mixed concrete association, fabeko. there is also the epd tool developed by basf for manufacturers to assess concretes [26] and other manufacturer specific epd tools such as that for tarmac [27]. therefore, when selecting a concrete, we recommend that specifiers do the following: (1) consider the range of embodied carbon impacts shown in figure 3; (2) ask local concrete producers if they are able to provide information on the embodied carbon (carbon footprint or epd) for their concretes; (3) look for a producer able to provide a concrete at the lower end of the embodied carbon range for its strength; and (4) make sure that any impacts from extended transport distances do not outweigh the benefits of reduced embodied carbon in production provided in module a4 of epd. 4.3. in building lca during early design stages, we recommend using regional generic data for concrete where available, as at this stage in the design this is the type of data recommended by en 15978 [28]. if generic data is not available for a particular region, then we recommend identifying a region with similar technology for cement and concrete production and using figure 2 to pick an appropriate embodied carbon value (from an industry generic epd if available, or from the median of the range otherwise). ganassali recommends the median as it is not sensitive to the outliers in a sample composed of a small number of data [16]. at later stages of the design, it is recommended to use specific data, e.g., from manufacturer specific epd based on the products you have chosen to use in the building (see section on material selection). 4.4. in benchmarking benchmarks can be provided for a product generally (e.g., cement or concrete) or for a specific product, such as cem i cement or c30 concrete. specifying benchmarks at the more specific level will ensure that the products meeting the benchmark are not just lower carbon products, but products which have lower carbon impacts than other products with similar functionality. for concretes, it is important that the functionality is considered in defining the benchmark. concretes with lower embodied carbon per m3 may have lower compressive strength, and this may lead to a requirement to use more concrete which could have an adverse environmental impact. for situations where compressive strength is relevant, we recommend using carbon intensity (co2eq/m3.mpa) as shown in figure 4 to set benchmarks. 22 vol. 33/2022 analysis of concrete and cement epd figure 1. graph showing relationship of embodied carbon for cements by clinker content, source [7]. figure 2. embodied carbon per m3 concrete by compressive strength, shown for each country, source [7]. 23 jane anderson, alice moncaster acta polytechnica ctu proceedings figure 3. box and whisker graph showing mean, upper and lower quartiles of embodied carbon for ready-mix concretes from concrete epd, by 28-day strength, source [7]. figure 4. box and whisker graph showing mean, upper and lower quartiles for carbon intensity ready-mix concretes, source [7]. 24 vol. 33/2022 analysis of concrete and cement epd in setting benchmarks, the geographical and related technological situation should also be considered ideally a regional benchmark should ensure that at least some of the production achieves the benchmark, whilst also stretching producers adopting "business as usual". 4.5. in setting reduction targets the lower quartile ranges in figure 3 and figure 4 show the best performance shown by 25% and 50% of products with existing epd, and therefore what should be achievable in setting long-term reduction targets for the majority of the market. this is also important information for manufacturers who wish to stay competitive in the carbon-aware market. 5. conclusion this paper has demonstrated how the analysis of existing epd for cement and concrete can provide multiple useful information for verifiers, specifiers, designers, building assessors and manufacturers, including for epd verification, material selection, building lca, benchmarking and target setting. a similar detailed analysis of many other products groups from the construction sector, particularly other materials and products with major impacts, would be useful for the construction industry as a whole. acknowledgements the authors are very grateful to the whitbybird foundation for funding j. a.’s initial work on the original paper [7]; and to the open-oxford-cambridge arts and humanities research council (ahrc grant number ah/r012709/1) doctoral training partnership for funding the continued project. funding for a. m. came from the school of engineering and innovation at the open university references [1] international energy agency (iea). global status report, 2018. https://webstore.iea.org/download/direct/2408? filename=2018_global_status_report.pdf. 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[28] cen/tc 350. en 15978:2011 sustainability of construction works assessment of environmental performance of buildings calculation method. brussels, 2011. 26 https://doi.org/10.1007/s11367-015-0926-3 https://doi.org/10.1007/978-3-319-66981-6_55 https://www.eco-platform.org/who-is-participating.html https://docs.google.com/spreadsheets/d/1ls7ukmug1cawnmghkiqivgcgehqoeiciih5t95inzy8/pubhtml https://doi.org/10.1680/adcr.13.00013 https://www.fedbeton.be/docs/technique/651_12_fr.pdf https://quantis-intl.com/csi-epd-tool-cement/ https://www.basf.com/us/en/media/news-releases/2014/01/p-13-478.html http://sustainability-report.tarmac.com/wp-content/uploads/2016/08/declaring-more-product-information.pdf https://doi.org/10.14311/app.2022.33.0404 acta polytechnica ctu proceedings 33:404–410, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague thin concrete overlays wit carbon reinforcement on deteriorated concrete pavements julia neumann∗, rolf breitenbücher ruhr university bochum, department of civil and environmental engineering, institute for building materials, universitätsstrasse 150, 44801 bochum, germany ∗ corresponding author: julia.neumann-a4n@rub.de abstract. in many countries, concrete pavements are normally built as jointed plain concrete pavements. due to a lack of alternatives, repairing deteriorated concrete pavements usually requires the replacement of the complete pavement structure and maintaining the joints, which is labourand resourceintensive. to increase the durability of repairs and to save resources concrete overlays with carbon reinforcement are developed. by the application of non-corrosive carbon-textile reinforcement cracks might be distributed so fine, that such an overlay can be executed jointless, unlike in previous repair methods. for a durable repair the bond behaviour between the retained concrete and the overlay as well as between the overlay-concrete and the textile reinforcement have to be considered. in this paper, the basic principles and feasibility of such a repair method are examined. on the one hand, the decisive influencing variables and parameters such as bond behaviour and cracking behaviour within the overlay are pointed out and discussed. on the other hand, the performed special lab tests will be presented. these tests include cyclic loadings on large-scale beams with integrated overlays of such types, evaluating the bond behaviour and the durability after a few millions of load cycles. furthermore, the crack formation in the overlay is determined by means of tensile and flexural tensile strength tests. keywords: carbon reinforced concrete, carbon reinforcement, jointed plain concrete pavements, maintenance, thin concrete overlays. 1. introduction in general, concrete pavements are exposed to considerable thermal and hygric stresses in addition to mechanical stresses. seasonal and diurnal temperature and moisture fluctuations cause temperature or moisture gradients in the cross section of the pavement structure thus resulting in constraint stresses and deformations. these in turn are expressed as tensile or compressive stresses on the topside of the pavement [1, 2] as well as over its entire cross section. overriding traffic, as a mechanical cyclic stress, causes bending stresses as well as degradation of the pavement concrete which can lead to material fatigue, as recent studies indicate [3]. to cope with these stresses, different construction methods are suitable for concrete pavements. in germany, the construction of jointed plain concrete pavements (jpcp) has historically prevailed [4] and is regulated in the german guideline "ztv betonstb 07" [5]. these structures are divided into slabs by joints. on the one hand, contraction joints are cut into the concrete as predetermined breaking points when the slab is still young, in order to avoid wild cracking and to enable the compensation of length changes. on the other hand, longitudinal joints are created due to production flow. in order to allow movement in the longitudinal direction and at the same time ensure the transfer of lateral forces between the pavement slabs and thus secure their height relative to each other, dowel bars are inserted into the green concrete in the contraction joints. the slabs are prevented from moving by tie bars in the longitudinal joints. to prevent dirt and water from penetrating into the substructure, the joints are sealed with suitable joint sealants or profiles. these must be designed in such a way that the longitudinal deformations of the slabs can continue to occur undisturbed. in order to guarantee the functionality of the joint seals and due to their limited service life of approx. 6 to 10 years, they require regular maintenance and servicing [6]. in addition to considerable costs from the maintenance work itself, this causes also massive traffic disruptions. 2. approach for a jointless repair concept regardless of the chosen construction method, concrete pavements will have to be repaired sooner or later. different repair methods are used depending on the nature and cause of the damage [7]. lacking marketable alternatives, in germany the entire superstructure is usually removed and rebuilt, if the serviceability and load-bearing capacity of concrete pavements cannot be restored by conventional repair measures due to damage or external factors, which 404 https://doi.org/10.14311/app.2022.33.0404 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 thin concrete overlays with carbon reinforcement figure 1. schematic construction of a jointless carbon textile reinforced concrete pavement (ctrcp) overlay on a deteriorated jpcp bottom-layer (jointed). materials basic composition: exposed aggregate concrete 8 mm cement cem i 42,5 n (sd) 430 kg/m3 w/c-ratio 0.42 sand 0 − 2 mm 487 kg/m3 basalt 2 − 8 mm 1310 kg/m3 grading curve − a/b 8 superplasticizer pce based 0.60 % by mass of cement air-entraining agent highly concentrated 0.15 % by mass of cement table 1. concrete mix design. is neither ecologically nor economically sensible [8]. abroad, on the other hand, there are various methods for large-scale repairs. in general, the damaged concrete layers are milled off and then reprofiled by use of unreinforced or reinforced overlays or precast elements [9–11]. bute these repair methods have limitations. for unreinforced overlays, the joints must be mirrored into the repair layer thus the scope of future maintenance cannot be reduced. precast elements, on the other hand, are mainly suitable for the repair of smaller areas, such as bus stops. reinforced overlays must be installed in relatively thick layers to protect the corrosion-prone reinforcement. the repair concept examined here is intended to avoid these disadvantages. after removing the damaged concrete layer (e.g. by milling or ultra-high pressure water jetting), it is planned to install a thin, jointless and continuously reinforced overlay. in order to achieve the planned layer thicknesses of only 30 to 70 mm, corrosion-resistant carbon reinforcement is used (see figure 1). the basic idea is to transfer the singular relatively large deformations in the joint area of the retained concrete, which are mainly caused by thermal and hygric influences, into multiple deformations in the form of fine cracks in the carbon textile reinforced concrete. 3. laboratory tests the laboratory tests were carried out on large scale beams to gain knowledge about the material behaviour of the carbon concrete overlay in combination with the retained concrete. for this purpose, the bond behaviour as well as the crack formation behaviour of the composite system of old concrete and carbon concrete were investigated in flexural tensile tests under static and cyclic load. 3.1. test specimen the tests were carried out on large beams consisting of a typical pavement concrete mix according to the german guideline "tl beton-stb 07" [12] (c30/37, xf4 or xm2 exposure class, maximum grain size 22 mm), which represents the deteriorated retained concrete, and a carbon textile reinforced concrete overlay (ctrc overlay) (consistency f4, air void content in fresh concrete 4.5 % by volume), which is realized in a concrete mix typically for exposed aggregate concrete with a maximum grain size of 8 mm (see table 1). the decisive advantage that carbon reinforcement offers for use in thin overlays is its corrosion resistance that allows small concrete covers of approx. 5 − 10 mm. in addition, the small diameters of the carbon rovings combined with its high tensile strength (2000 − 6000 mpa [13]) also contribute to thin overlays. here, a carbon mat reinforcement was used [14]. it was equipped with a factory applied polystyrene coat405 julia neumann, rolf breitenbücher acta polytechnica ctu proceedings overlay thickness number of textile layers average of type of failure number of cracks crack spacing [mm] [−] [−] [mm] [−] 70 2 2.5 65.2 delamination of reinforcement /fracture of test specimen 3 5.0 50.0 delamination of reinforcement 50 1 2.0 75.0 delamination of reinforcement / fracture of test specimen2 3.5 79.03 4.0 30.0 30 1 2.5 112.5 delamination of reinforcement 2 2.0 150.0 delamination of reinforcement /fracture of test specimen table 2. average number of cracks and crack spacing above bond breaker. figure 2. textile reinforcement made of carbon rovings. ing which is intended to improve the bond between reinforcement and concrete. the textile (see figure 2) had a centre distance of the rovings of 12.7 mm in longitudinal (cross section 141.02 mm2/m) and of 16.0 mm in transverse (28.02 mm2/m) direction, resulting in a mesh size of 9 mm × 14 mm. due to the maximum grain size of the aggregate of 8 mm complete enclosure of the reinforcement mats by the concrete without major air inclusions was only possible if the concrete was placed and compacted in layers between the reinforcement mats. the test specimens obtained had a length of 1800 mm, width of 250 mm and 500 mm and varied in height between 180 mm and 270 mm, depending on the planned tests and configurations of the ctrc overlay (see figure 3). the overlays thickness varied between 30, 50 and 70 mm and it was reinforced with 1 to 3 layers of carbon reinforcement. two large beams were investigated for each parameter combination. before the overlay was applied the retained concrete (150 − 240 mm in height) were roughened by high-pressure water jetting and afterwards the surface was cleaned. the contraction joint between two pavement slabs was simulated by placing a 3 mm thick pvc foil for the cyclic tests and a recess without filler material for the static tests. in addition to this, two dowel bars (250 mm distance, 500 mm length and 25 mm diameter) were placed mid-height of the specimen for the cyclic tests. a bond breaker in form of a pe construction foil of 150 mm width was placed above the contraction joint in the bond joint between retained concrete and ctrc overlay to allow a free expansion length and thus promote the formation of fine cracks in the overlay and prevent reflection cracking [15]. 3.2. test program the following described tests and results represent an extract from the research project. a comprehensive presentation can be found in the final report of the project. 3.2.1. flexural tensile tests the aim of the flexural tensile tests under static load was to gain knowledge about the crack development of the ctrc overlay. for this purpose, the test specimens were subjected to 4-point flexural tensile tests at a loading speed of 1 mm/min according to figure 3. in addition to the documentation of the position, size and order of occurrence of the individual cracks, the strains that occurred during the tests were recorded simultaneously to the applied force. independent of the configuration of the carbon concrete layer, almost all test specimens showed similar crack patterns to figure 4. shear cracks occurred, running from the supports to the load inducing points 406 vol. 33/2022 thin concrete overlays with carbon reinforcement figure 3. specimens for flexural tensile tests. 1: bearing, 2: dowel, 3: simulated contraction joint, 4: bond breaker, 5: dowel bar. figure 4. example of crack pattern. 1: bearings, 2: carbon textile reinforcement layers, 3: joint, 4: bond breaker, 5: shear cracks, 6: horizontal cracks, 7: delamination. 407 julia neumann, rolf breitenbücher acta polytechnica ctu proceedings figure 5. example of the development of the relative cyclic young’s modulus over 2 million load cycles on a 30 mm thick cracked ctrc overlay (2 reinforcement layers, with bond breaker) as a indicator for micro-cracking. (1). in between, horizontal cracks were formed in the areas under compressive stress (2). in many cases, delamination occurred starting from the shear cracks (3), which spread in the direction of the supports and, until failure, also in the direction of the beam centre. depending on the configuration of the carbon concrete layer, different crack patterns formed between and outside the described cracks (4). an influence of the dowels in the bond joint on the resulting crack patterns could not be determined so far. in addition, more cracks tended to occur in the joint area if a bond breaker was used. then up to seven cracks with a distance of 30 − 150 mm could be observed. without a bond breaker only up to 3 cracks with a distance of 75 − 150 mm could be observed. furthermore, the number of cracks tended to increase with the number of reinforcement layers, although this could not be observed for a carbon concrete layer thicknesses of 30 mm (see table 2). 3.2.2. cyclic flexural tensile tests the bond behaviour of retained and carbon concrete as well as carbon reinforcement and concrete was investigated by flexural tensile tests under cyclic load. here, too, a 4-point flexural tensile test was used to simulate the loads occurring in situ in the joint area. the combination of the load cases "traffic load + rapid cooling" leads to the highest stresses and differences in stress and is therefore used as the design load for the slab edge. the maximum constraint stress of σo = 3.40 mpa results from the superimposition of the two load cases traffic load σqr = 1.2 mpa and rapid cooling. with the relevant temperature gradient of 0.4 k/cm, the latter generates a reduced buckling stress σw ≈ 2.2 mpa as a tensile stress, which represents the lowest stress in the cyclic flexural tests [16]. based on the determined flexural tensile strength of the unreinforced carbon concrete of 5.610 mpa this results in a degree of utilisation of σo 6/fc, t ≈ 0.61. more detailed information on the resulting constraint stresses and the determination of the decisive design load can be found in [16]. the large beams were installed in the hydraulically controlled test stand in such a way that the ctrc overlay was under tensile stress. the test were carried out with a loading frequency of 5 hz. during the test, the concrete surfaces were examined at regular intervals for crack formation and delamination and the change in the relative cyclic young’s modulus, as an indicator of micro-cracking, was determined by means of ultrasonic measurement. at the same time, the strain above the contraction joint was continuously recorded by means of a strain gauge glued to the centre of the carbon concrete surface. after two million load cycles, drill cores were taken from loaded and unloaded beam areas for further laboratory tests, the results of which are not shown here. in addition to shear tests and centric tensile tests to 408 vol. 33/2022 thin concrete overlays with carbon reinforcement determine the effective composite forces, these also included leakage tests, as well as investigations of the freeze-thaw resistance in the loaded joint area. in the tests under cyclic load carried out so far, cracks were observed on some large beams over the entire beam width. after only 10,000 to 50,000 cycles, a significant drop in the relative cyclic young’s modulus of 53 − 79 % (to 21 − 47 % of the zero value) was recorded in these cases at the measuring section concerned (example see figure 5). these cracks were macroscopically visible from 50,000 load cycles on, whereby they showed a crack width of approx. 0.1 mm. while in some cases one crack occurred in the middle of the beam, in the remaining cases two further cracks occurred directly at the edges of the bond breaker. cracks tended to occur on beams with thinner carbon concrete layers (30 − 50 mm) and 1 to 2 reinforcement layers, although the crack pattern was not always the same on beams with the same configuration of the overlay. spalling did not occur on the crack flanks. if no cracks formed, the relative cyclic young’s modulus of elasticity dropped by 9−19 % (on average 11.2 %) after 2 million load cycles. the values fluctuated over the length of the beam, but no clear tendencies were discernible. in previous research projects, 10−15 % decrease in the relative cyclic young’s modulus was observed for jpcp [3]. based on these results, structural damage could be determined to an extent which is also to be expected in non-reinforced concrete pavements. in the cracked ctrc overlays similar damage was observed outside the cracks. furthermore, no delamination of concrete and reinforcement was observed. tensile tests carried out on drill cores from loaded and unloaded beam areas also only showed failure in the bonded joint of retained and ctrc overlay. 4. final discussion and conclusion in the flexural tensile tests at 20 ◦c on large concrete beams with ctrcp overlays, it was shown that the contraction joint can be bridged without reflection cracking and at the same time can be transferred into several fine cracks in the overlay. although the number of cracks tends to increase with used reinforcement layers, fewer cracks have formed than desirable. thus a wider bond breaker seems useful and should be investigated. the failure types to be observed are the fracture of the test specimen with associated tearing of the reinforcement and the delamination of the overlay concrete from the reinforcement. no test specimen has shown a failure in the bond joint between retained concrete and overlay. under service load, cracks could be observed on some beams directly above the contraction joint and the edges of the bond breaker. up to 2 million load cycles they remained relatively constant in width as soon as they were macroscopically visible. further investigations of water absorption and frost resistance are being carried out to investigate the effect of the cracks on the durability of the overlay. if no cracks occurred, the structural damage was comparable to the damage of concrete pavements under cyclic load observed in previous projects. to what extent these damages influence the durability of the ctrc overlay should be investigated in the future. furthermore, the behaviour of ctrc overlay and retained concrete under pressure and centric tensile load should be investigated. additionally it is necessary to research the functionality of the repair method in the form of test tracks. acknowledgements the authors would like to thank the federal ministry of education and research (bmbf), germany, represented by forschungszentrum jülich and the committee of the collaborative research project "c3 carbon concrete composite" for their funding of this research project. further, the authors would like to acknowledge the fruitful cooperation with the institute of concrete structures, technische universität dresden. their partner project has added valuable aspects to the findings of this 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[16] c. sievering. dauerhaftigkeit von betonfahrbahndecken unter besonderer berücksichtigung des externen alkalieintrags, bochum: institute for building materials, ruhr university bochum, ph.d. thesis, 2012. 410 https://www.pci.org/pci_docs/publications/pci%20journal/2013/winter/precast%20concrete%20pavements%20-%20technology%20overview%20and%20technical%20considerations.pdf http://dx.doi.org/10.18419/opus-192 https://doi.org/10.1515/rbm-2011-6452 https://doi.org/10.14311/app.2022.33.0631 acta polytechnica ctu proceedings 33:631–637, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague quantifying uncertainties in the sustainability evaluation of concrete materials considering regional characteristics in japan nicole alexis kwan viosa, ∗, michael ward henryb, joel galupo oponc a hokkaido university, graduate school of engineering, kita 13 nishi 8, kita-ku, sapporo 060-8628 japan b shibaura institute of technology, department of civil engineering, toyosu 3-7-5, koto-ku, tokyo 135-8548 japan c msu-iligan institute of technology, department of civil engineering, faculty member, iligan city 9200 philippines ∗ corresponding author: nicolealexisv@gmail.com abstract. in the evaluation of concrete sustainability, what constitutes "sustainable" to one region may vary from another. this often leads to methodological forms of uncertainties that makes the evaluation process more complex. as such, this paper aims to quantify the effect of uncertainties in the regional context on the sustainability evaluation of concrete materials. this is carried out by quantifying the regional context through establishing a weighting scheme and then integrating the obtained weights into the sustainability analysis of concrete materials in tandem with uncertainty analysis. japan is used as a case study because although it relatively appears as a homogeneous country, its prefectures possess unique characteristics that may make the sustainability evaluation of concrete materials vary across prefectures. cluster analysis is carried out in the 47 prefectures of japan using a set of regional context indicators. five clusters are identified with varying characteristics and these are translated into different weighting schemes. the established weights are used in the sustainability evaluation of concrete materials using multi-criteria decision-making analysis. the results showed that one mix is the most sustainable for four of the clusters and a different mix is the most sustainable for the remaining cluster. when uncertainty analysis is conducted, the effect of the weights in the sustainability evaluation is explained by examining the average scores of the concrete mixes and the variance of the scores across the five clusters. this investigation facilitated the understanding of how regional differences and the uncertainties associated with it impact the evaluation of concrete sustainability. keywords: regional characteristics, sustainable concrete, uncertainty analysis. 1. introduction sustainable development, commonly defined as "the development that meets the needs of the present without compromising the ability of future generations to meet their own needs" [1], has its implementation dependent on individual governments. each country has varying context on this concept making the means of attaining sustainability different among countries [2]. with the growing awareness of sustainable development, the concrete industry is taking efforts to participate in the practice of sustainability. similar to sustainable development, the implementation of sustainability in concrete also varies by region because concrete is a regional context material itself. henry and opon [3] introduced a means for quantitatively analyzing the regional context of sustainability and linking this regionality to the sustainability evaluation of concrete materials. however, methodological forms of uncertainties exist in this evaluation which make the evaluation process more complex. as such, it is the goal of this paper to quantify the effect of uncertainties in the regional context on the sustainability evaluation of concrete materials. japan was used as a case study for this analysis because although japan appears to be relatively homogeneous as a country, it has 47 prefectures having unique characteristics which may make the sustainability evaluation of concrete materials vary across prefectures. 2. methodology the analytical flow adopted in this study is summarized in figure 1. initially, a set of quantitative regional context indicators was constructed, gathered for all prefectures and processed accordingly. then, an agglomerative hierarchical clustering method was used to explain the regional context by identifying groups of prefectures exhibiting similar characteristics. these clusters form the basis of weighting schemes which represent regionality and the obtained weights were used as inputs in the concrete sustainability evaluation of a set of concrete mix utilizing multicriteria decision-making analysis. finally, the 631 https://doi.org/10.14311/app.2022.33.0631 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en n. a. vios, m. w. henry, j. g. opon acta polytechnica ctu proceedings figure 1. analytical flow. figure 2. sustainability evaluation framework using mca. range of z-score z≤-3 -3<z≤-2 -2<z≤-1 -1<z≤-0.5 -0.5<z<0.5 0.5≤z<1 1≤z<2 2≤z<3 3≤z relative intensity (−) 1/9 1/5 1/3 1/2 1 2 3 5 9 relative intensity (+) 9 5 3 2 1 1/2 1/3 1/5 1/9 table 1. conversion of in-cluster mean z-scores to relative intensities for calculation of weights. effect of methodological uncertainties in the regional context of concrete sustainability evaluation was investigated using uncertainty analysis to understand how these uncertainties influence concrete sustainability evaluation results. 2.1. cluster analysis methodology and extraction of weights in accordance with the guidelines provided by kassambara [4] for conducting cluster analysis, the agglomerative hierarchical clustering method was adopted using the open source statistical analysis software r [5]. in this approach, the prefectural data was first standardized to a mean of zero and a standard deviation of one converting them to z-scores. after which, the distance matrix which summarizes the distance between any two prefectures in an n dimensional space (where n is the number of indicators) was calculated using the euclidean distance method. a hierarchical tree was created using a linkage function to group prefectures based on their distance information. finally, a dendrogram, which is a visualization of the cluster analysis results, was produced and examined for analyzing the regionality of concrete sustainability. the clusters were the basis in the extraction of weights. weighting of indicators is an important area in sustainability evaluation as they reflect the significance of different dimensions in their contributions to the sustainability performance of a system [6]. a weighting scheme based on the analytic hierarchy process (ahp) approach was used to reflect regionality in the evaluation of concrete sustainability similar to the method taken by henry and opon [3]. ahp is a multi-criteria decision-making tool that structures the decision-making process as a hierarchy composed of quantifiable criteria and their relations and alternatives towards a goal [7]. weight determination is dependent upon pairwise comparison of indicators and since pairwise comparison is not applicable to negative or zero values, the z-scores were converted to relative intensities following the scale shown in table 1 depending on the loading direction of the indicator as shown in table 5. 2.2. concrete sustainability evaluation methodology after obtaining the different weights, sustainability evaluation of the concrete mixes was conducted using multi-criteria analysis (mca) shown in figure 2. each step in the framework is explained in the following sections. 632 vol. 33/2022 sustainability evaluation of concrete materials in japan label data source description (units) res1 a total cement consumption in concrete per capita (100kg/capita) res2 a total aggregate consumption in concrete per capita (100kg/capita) res3 b total water withdrawal per capita (m3/person) res4 c usage rate of slag blended cements (0% of total cement consumption) res5 c usage rate of fly ash blended cements (0% of total cement consumption) wst1 d total concrete construction waste generated per capita (100kg/person) eco1 d total construction investment per capita (1000yen/person) env1 e total co2 emissions per capita (tons/person) env2 e total energy consumption per capita (gj/person) data source: a: ministry of economy, trade and industry b: e-stat c: japan cement association d: ministry of land, infrastructure and transportation e: ministry of the environment table 2. indicators for regional context analysis data from 2010. 2.2.1. indicator selection indicators are selected based on their relevance to what is being measured. they are the fundamental structural basis on which the sustainability of a concrete material can be judged [8]. 2.2.2. normalization the indicators are expressed in different units. in order for them to work together in an aggregated composite, normalization is necessary. several methods are available for this purpose but for this paper standardization using z-scores was used. 2.2.3. weighting weighting schemes can sometimes increase or eradicate trade-offs between indicators, which makes weighting a controversial issue in concrete, as current codes are silent about it and still emphasize the hierarchy of mechanical performance [9]. due to this subjectivity, clarifying how much of an impact the weights have in the concrete sustainability evaluation is demonstrated in this analysis. 2.2.4. aggregation geometric aggregation (g) was chosen as the aggregation method in this paper as it involves aggregation by geometric mean using the product values [8] and is appropriate when less compensability is desired [10]. equation 1 shows how this is calculated where w is the weight assigned to each indicator and sci is the indicator normalized value. g = n! i=1 = (sci) wi (1) finally, the aggregated values were standardized using z-scores followed by a transformation to t-scores converting the results to a scale of 0 to 100 assuming a mean of 50 and herein referred to as sustainability score (ss). 3. results and discussions 3.1. regional context indicators for quantifying the regional context of concrete sustainability, a set of 9 indicators was constructed as summarized in table 2 spanning four categories. the resource consumption indicators (total cement, aggregate and water) and the waste generation indicator (total concrete construction waste) focus on the flow of materials from the beginning to end of its life cycle. the distribution of steel manufacturing and coal power generation within japan affects the usage of blended cements leading to the inclusion of two additional resource consumption indicators (usage rate of slag blended and fly ash blended cements). finally, economic and environmental aspects of sustainability were represented with three indicators measuring total construction investment, total co2 emissions and total energy consumption. normalized values (e.g. per capita) were utilized to facilitate comparison of variables that scale with prefectural size. 3.2. clustering analysis with the regional context indicators, cluster analysis was carried out to identify groups of prefectures exhibiting similar characteristics. the linkage function used was the average linkage function because it had the highest correlation coefficient of 0.86 between the euclidean distance matrix of the standardized data and the cophenetic distance matrix among the several linkage functions tested. the dendrogram produced by the clustering result is shown in figure 3. the authors set the number of clusters to five by cutting the tree at the height shown to sufficiently represent regionality without overly simplifying nor overly complicating the analysis. clusters are numbered from one to five from the left. the in-cluster mean z-scores of the five clusters for each regional context indicator are summarized in table 3. clusters 1 and 4 are single-membership clusters 633 n. a. vios, m. w. henry, j. g. opon acta polytechnica ctu proceedings figure 3. dendrogram of agglomerative hierarchical clustering results. cluster res1 res2 res3 res4 res5 wst1 eco1 env1 env2 1 0.38 0.45 -0.52 -1.44 6.04 0.31 1.98 -0.18 -0.16 2 0.38 -0.14 -0.17 0.67 -0.24 0.84 -0.39 3.17 3.19 3 -0.28 -0.26 0.04 0.02 -0.12 -0.27 -0.22 -0.19 -0.18 4 1.34 1.02 -1.58 1.29 -0.24 3.69 2.00 -0.37 -0.55 5 1.91 2.20 0.32 -0.67 -0.15 0.92 1.34 -0.42 -0.47 table 3. in-cluster mean z-scores for regional context indicators by cluster. containing aomori and shimane, respectively, that is the in-cluster mean z-scores of these clusters are the z-scores of the prefecture themselves. the in-cluster mean z-scores explain the distinctive characteristics of each cluster and these were used as inputs in calculating the weights. 3.3. extraction of weights the calculated weights are shown in table 4 and it can be deduced from these results that different clusters put different weights on the indicators. as these values were normalized to have a total sum of 1, it is easier to compare than the in-cluster mean z-scores. cluster 1 consisting aomori puts a very high weight on res5 derived from the high value of the prefecture’s fly ash usage rate. cluster 2 composed of oita, okayama and yamaguchi puts the same emphasis on env1 and env2 attributed to the high value of its co2 emissions and energy consumption. cluster 3 containing 38 clusters have equal weighting for all the indicators and can be said to represent the average conditions in japan. cluster 4 with only shimane has its greatest weight on wst1 associated with its high concrete construction waste. finally, cluster 5 comprising okinawa, toyama, fukui and yamanashi has a number of indicators with high weight including res1, eco1 and especially res2 related to its high total aggregate consumption. these weights are parallel with the in-cluster mean z-scores of the clusters. 3.4. concrete sustainability evaluation a set of 9 indicators for concrete sustainability evaluation was constructed as shown in table 5. each concrete sustainability indicator has its corresponding regional context indicator. the table shows the loading direction of each indicator which indicates the behaviour that contributes to increased sustainability and is employed in the calculation of weights. for example, decreasing the consumption of raw materials improves the sustainability of a concrete mix. a test set of three concrete mixes was drawn from noguchi et.al. [11] and the mix proportions are shown in table 6. to render consistency in the analysis, mixes with 28th day compressive strengths of approximately 30mpa which are generally used for civil engineering infrastructures were chosen. three new alternatives were created containing high-grade recycled aggregates are also shown in the table. highgrade recycled aggregates were assumed to have the exact same properties and cost as normal aggregates 634 vol. 33/2022 sustainability evaluation of concrete materials in japan cluster weight (w) res1 res2 res3 res4 res5 wst1 eco1 env1 env2 1 w1 0.05 0.05 0.02 0.15 0.44 0.05 0.15 0.05 0.05 2 w2 0.04 0.04 0.04 0.07 0.04 0.07 0.04 0.33 0.33 3 w3 0.11 0.11 0.11 0.11 0.11 0.11 0.11 0.11 0.11 4 w4 0.12 0.12 0.01 0.12 0.04 0.35 0.19 0.04 0.02 5 w5 0.16 0.26 0.05 0.11 0.05 0.11 0.16 0.05 0.05 table 4. indicator weights for each cluster. label indicator units loading related regionalcontext indicator cem cement content kg/m3 − res1 agg natural aggregatecontent kg/m 3 − res2 wat water content kg/m3 − res3 bfs blast furnaceslag content kg/m 3 + res4 fa fly ash content kg/m3 + res5 rca recycled aggregatecontent kg/m 3 + wst1 cst construction materialscost yen/m 3 − eco1 co2 co2 emissions kg−co2/m3 − env1 ener input energy mj/m3 − env2 table 5. indicators for concrete sustainability evaluation. mix 28th day f ′c w/b w c bs fa s na ra (mpa) (−) (kg) (kg) (kg) (kg) (kg) (kg) (kg) ref 30.57 0.57 184 325 0 0 783 1063 0 a 30.57 0.57 184 325 0 0 783 0 1063 b1 30.44 0.58 192 200 133 0 806.2 965.4 0 b2 30.44 0.58 192 200 133 0 806.2 0 965.4 f1 30.23 0.56 186.7 238.1 0 94.1 847 949.9 0 f2 30.23 0.56 186.7 238.1 0 94.1 847 0 949.9 table 6. concrete mix proportions per m3. for simplicity. thus, concrete with high-grade recycled aggregates have the same properties as mixes with normal aggregates, but the inventory data shows that they require more energy to process to go back to the original state which is reflected by higher co2 emissions and energy consumption [12]. table 7 shows the calculated indicator values for each mix wherein the six constituent material indicators were all determined directly from the mix proportions. due to the lack of data on the regional variation of the costs of constituent materials, this analysis uses a single value per mix obtained from henry and opon [3] and calculated to get the unit cost per mix. lastly, the two environmental impact indicators were computed using japanese inventory data [13]. the concrete sustainability evaluation results using mca are shown in figure 4. each cluster’s results represent one implementation of mca. looking at each cluster, varying weighting schemes lead to varying sustainability scores for the different mixes. for clusters 1, 3, 4 and 5, the most sustainable concrete material is only one mix (f2). while for cluster 2 the most sustainable concrete material is a different mix (b1) whose selection is heavily supported by the high weight placed on co2 and ener indicators by this cluster and among the concrete mixes, mix b1 has the lowest co2 emission and energy consumption values. despite the same most sustainable concrete material for four of the clusters, the ranks of the other mixes change. but, a high level of confidence can be presumed on mix f2 considering a general choice for the most sustainable concrete material for the entire country. 635 n. a. vios, m. w. henry, j. g. opon acta polytechnica ctu proceedings mix cem agg wat bfs fa rca cst co2 ener ref 325 1063 184 0 0 0 5,775.04 255.12 1213 a 325 0 184 0 0 1063 5,775.04 270.86 1563.8 b1 200 965.4 192 133 0 0 5,679.39 162.63 861.84 b2 200 0 192 133 0 965.4 5,679.39 176.92 1180.4 f1 238.1 949.9 186.7 0 94.1 0 5,265.96 190.22 956.62 f2 238.1 0 186.7 0 94.1 949.9 5,265.96 204.28 1270.1 table 7. concrete mix characteristics (units are shown in table 5). figure 4. sustainability scores of concrete mixes by cluster. mix average ss rankaveragess variance rankvariance ref 36.00 6 7.27 1 a 41.08 5 29.11 5 b1 50.55 4 32.62 6 b2 57.09 2 18.46 2 f1 54.41 3 27.18 4 f2 60.87 1 21.61 3 table 8. descriptive statistics of uncertainty analysis for each concrete mix. 3.5. uncertainty analysis the subject of uncertainty propagation in concrete material sustainability evaluation that arises from weighting schemes is discussed in this section (i.e. how much the weights impact the sustainability evaluation results). the method undertaken for this analysis is where the ss of each mix for all of the clusters were aggregated and the corresponding average and variance together with their individual ranks were computed as shown in table 8. when choosing a sustainable concrete material in a national level but at the same time being aware of the existence of the regional context, saying that mix f2 is the most sustainable concrete mix because it is ranked 1st by four out of five clusters has uncertainties associated with it. the uncertainty analysis is a way of quantitatively examining how much consideration of the regional context will affect the selection of a sustainable concrete mix. as an example, mix f2 was ranked 1st based on the average sustainability score as it showed to be the most sustainable concrete material for four of the clusters, but considering the variance of this selection, it ranked 3rd which can reasonably be identified as sustainable. however, when looking at mix b2 which ranked 2nd based on the average sustainability score, it showed a lower variance. this then is where the choice between choosing a moderately scored material with low variance versus a highly scored material with moderate variance comes in. generally, a mix with higher average sustainability score and lower variability is desired as this entails a more confident result. based on this idea and referring to figure 5, the bottom right mix is the most desirable decision and is identified to be mix f2. these results cannot be an output by conducting concrete sustainability evaluation alone. viewing the mixes in the perspective of uncertainty analysis gives evidence that even though considering regional context in concrete sustainability evaluation as represented by the weights, the most sustainable concrete material for a national level nonetheless shows to be the same material that displayed the most sustainable character in a cluster level. 636 vol. 33/2022 sustainability evaluation of concrete materials in japan figure 5. scatter plot of average ss vs variance. 4. conclusions in this paper, quantifying the regional context of concrete sustainability was demonstrated that lead to obtaining different weighting schemes. this then was integrated in the evaluation of concrete sustainability. however, uncertainty analysis in this evaluation suggests that those weights, while they can lead to different evaluations of concrete mixes at a mesoscale (cluster level), the effect of the weights potentially may be negligible in a macroscale (national level) considering that the assumptions made in the analysis holds the same and true. this is evidenced by the varied sustainability scores for each mix per cluster, but the rank of the most sustainable concrete material does not change for one mix except in one cluster. with this at hand, it can be concluded that the weights are not leading to a significant difference of the selection of the most sustainable concrete material. therefore, applying a simple weighting scheme, such as equal weighting, may be sufficient for the representative sustainability evaluation of concrete. references [1] g. h. brundtland. our common future-call for action. environmental conservation 14(4):291-4, 1987. https://doi.org/10.1017/s0376892900016805. [2] m. henry, y. kato. sustainable concrete in asia: approaches and barriers considering regional context. construction and building materials 67:399-404, 2014. https: //doi.org/10.1016/j.conbuildmat.2013.12.074. [3] m. henry, j. opon. regional context indicators for the evaluation of concrete sustainability: an analysis at the prefectural level in japan. the 8th international conference of asian concrete federation, china, 2018. [4] a. kassambara. practical guide to cluster analysis in r, edition 1, sthda.com, 2017. [5] r core team. a language and environment for statistical computing r foundation for statistical computing vienna, 2018. http://www.r-project.org/. [6] x. gan, i. c. fernandez, j. guo, et al. when to use what: methods for weighting and aggregating sustainability indicators. ecological indicators 81:491-502, 2017. https://doi.org/10.1016/j.ecolind.2017.05.068. [7] t. l. saaty. decision making for leaders rws publications, rws publications, pittsburgh, 2008. [8] m. j. burgass, b. s. halpern, e. nicholson, et al. navigating uncertainty in environmental composite indicators. ecological indicators 75:268-78, 2017. https://doi.org/10.1016/j.ecolind.2016.12.034. [9] j. opon, m. henry. understanding the propagation of uncertainty in concrete material sustainability evaluation the 8th international conference of asian concrete federation, china, 2018. [10] m. saisana, a. saltelli. rankings and ratings: instructions for use. hague journal on the rule of law 3(02):247-68, 2011. https://doi.org/10.1017/s1876404511200058. [11] t. noguchi, f. tomosawa, k. m. nemati, et al. a practical equation for elastic modulus of concrete. aci structural journal 106(5):690-96, 2009. https://www.researchgate.net/profile/kamran-nem ati/publication/279700549_a_practical_equation _for_elastic_modulus_of_concrete/links/5cbd7a 30a6fdcc1d49a5ea2c/a-practical-equation-for-e lastic-modulus-of-concrete.pdf. [12] m. w. henry. formation and evaluation of sustainable concrete based on social perspectives in the japanese concrete industry, 2010. [13] t. uomoto, c. hashimoto. recommendation for shotcreting (draft) part tunnels by natm by the japan society of civil engineers. shotcrete for underground support x, p. 14-29, 2006. https://doi.org/10.1061/40885(215)3. 637 https://doi.org/10.1017/s0376892900016805 https://doi.org/10.1016/j.conbuildmat.2013.12.074 http://www.r-project.org/ https://doi.org/10.1016/j.ecolind.2017.05.068 https://doi.org/10.1016/j.ecolind.2016.12.034 https://doi.org/10.1017/s1876404511200058 https://www.researchgate.net/profile/kamran-nemati/publication/279700549_a_practical_equation_for_elastic_modulus_of_concrete/links/5cbd7a30a6fdcc1d49a5ea2c/a-practical-equation-for-elastic-modulus-of-concrete.pdf https://doi.org/10.1061/40885(215)3 https://doi.org/10.14311/app.2022.33.0256 acta polytechnica ctu proceedings 33:256–262, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague chloride ion penetration behaviour in concrete containing an expansive additive and a calcium-aluminate-based additive shinya itoa, ∗, takeshi iyodab a denka co.ltd., 2209, omi, itoigawa-city, niigata 949-0393, japan b shibaura institute of technology, 3-7-5, toyosu, kotoku, tokyo 135-8548, japan ∗ corresponding author: shinya-ito@denka.co.jp abstract. the effect of the material characteristics on the infiltration behaviour of chloride ions in concrete containing a combination of an expansive additive and a calcium-aluminate-based additive has been investigated. the same level of salt resistance as that of blast furnace slug cement was exhibited even for a small addition amount. in addition, the permeation behaviour of chloride ions was largely influenced by the immobilisation capacity and pore network. when these changes were dominant, the results suggested that the characteristics based on the fick diffusion equations may not necessarily reproduce the actual permeation behaviour of chloride ions. keywords: chloride ion, chloride ion immobilized material, diffusion coefficient, expansive additive. 1. introduction in japan, there is a large amount of airborne salt from the ocean, and degradation of reinforced concrete structures because of salt damage has been reported nationwide owing to spraying of antifreeze agents in winter. salt damage is generally indicative of steel corrosion in concrete owing to infiltration of chloride ions. for general-purpose materials, blast furnace cement and common additives, such as fly ash, are used. in recent years, an additive has been developed in which cao · 2al2o (ca2) of calcium aluminate is admixed with cement to react with ca(oh)2(ch), which is a cement hydrate, to form a hydrate with an immobilisation function, and free chloride ions, which cause corrosion of steel, are chemically immobilised as friedel’s salt. it has been reported that the additive exhibits high salt damage resistance at low dosage [1, 2]. in addition, if excessive cracks occur in the concrete structure, chloride ions directly reach the reinforcing rebar, which increases the possibility of premature deterioration. therefore, it is important to suppress crack occurrence as a salt damage countermeasure. in addition, by incorporating an expansive additive in concretes, it is possible to suppress excessive cracking by introducing a shrinkage reduction effect or a chemical pre-stressing effect [3]. therefore, it can be expected that the external permeation path of chloride ions can be minimised by using ca2 and an expansive additive in combination, and the permeation rate of chloride ions can be suppressed by the immobilisation function even in the cured body. although the basic characteristics and salt damage resistance of concretes containing ca2 and an expansive additive in combination have been reported in previous studies, it is essential to establish a method to reflect the results to check the durability of concrete for practical use. for commons additives, such as blast furnace slag and fly ash, the characteristic values for durability inspection are clearly indicated in the concrete standard statement of the society of civil engineers based on a vast amount of data, but they cannot be directly applied to concrete containing ca2 or an expansive additive. in this study, we investigated the material characteristics of concretes containing ca2 and an expansive additive, and we then investigated the effect of the material characteristics on the chloride ion infiltration behaviour for durability examination. 2. materials and methods 2.1. composition of materials and concrete ordinary portland cement (opc) and blast furnace cement b (bb) were used as the cements. ca2 was prepared by calcining at 1750 to 1850 ◦c. a cao/al2o3 molar ratio of 0.5 was used with calcium carbonate and aluminium oxide as the raw materials. the obtained clinker was pulverised by gradual cooling. a lime-ettringite composite was used as the expansive additive. the chemical composition and densities of ca2 and the intumescent materials are given in table 1. sand (density 2.62 g/cm3) produced by kimitsu city, chiba prefecture was used as the fine aggregate, and limestone crushed stone (density 2.70 g/cm3) produced by tsukumi city, oita prefecture was used as the coarse aggregate. the concrete mix proportions are given in table 2. the cement, ca2, and the expansive additive were regarded as a binder (b, table 2). 256 https://doi.org/10.14311/app.2022.33.0256 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 chloride ion penetration behaviour in ca-al concrete chemical composition (%) density (g/cm3) cao sio2 al2o3 so3 fe2o3 opc 64.1 20.5 5.2 2.1 3.0 3.16 ca2 24.0 0.6 67.7 0.01 7.1 2.96 ex 70.6 1.0 7.2 18.5 0.8 3.10 table 1. chemical compositions and densities. unit weigth (kg/m3) no. w/b (%) s/a (%) w c ca2 ex s g n 55.0 48 170 309 − − 864 965 n + ca2 289 20 − 964 n + ca2 + ex 269 20 22 964 bb 310 − − 859 959 table 2. concrete mix proportion. 2.2. test materials and measurement methods 2.2.1. permeation characterisation of the chloride ions a salt water immersion test was performed to evaluate the permeation characteristics of chloride ions. in the salt water immersion test, a 100 mm × 100 mm × 400 mm prismatic specimen was cured in water at 20 ◦c for 28 days. the five surfaces other than the 100 mm × 400 mm surface were coated with epoxy resin and immersed in 10% aqueous nacl solution at 20 ◦c with one surface exposed to epoxy resin. at the specified age of the specimen, the depth from the exposed surface to the coloured part was measured by sequentially splitting and spraying silver nitrate solution (0.1 n) on the split surface. 2.2.2. chloride ion profile after 48 weeks aging of the immersion material, the test piece was cut at intervals of 1 cm in the depth direction from the permeation surface of the specimen. the total chloride ion content was analysed in accordance with jis a1154, and the chloride ion profile in the depth direction was constructed. the soluble chloride ions were extracted with warm water at 50 řc, and the amounts of free and fixed chloride ions were determined. 3. results and discussion 3.1. evaluation of the chloride ion permeability 3.1.1. salt immersion test the chloride penetration depths of the concrete specimens soaked in 10% strength nacl solution are shown in figure 1. the penetration depth of the reference sample n aged for 48 weeks was about 30 mm. the penetration depths of the n+ca2 and n+ca2+ex samples were about 20 and 16 mm, respectively, and figure 1. salt immersion test results. the penetration of chloride ions tended to be suppressed by the use of additives. in addition, the difference in the penetration depth compared with bb was several millimetres, and there was not a large difference in the penetration depth in the immersion period. it was presumed that this was because of the fact that hydrates with an immobilisation function, such as hydroculmite (hc), were produced by adding ca2 immobilised chloride ions as friedel’s salts, thereby suppressing penetration of chloride ions into the concrete. when the expansive additive was used in combination with ca2, the penetration depth of chloride ions was equal to or less than that of the ca2 sample. in addition, if the expansion ratio was within the range of the expansion ratio of shrinkagecompensating concrete, there was no effect on the immobilisation capability of chloride ions of the ca2 sample, and a synergistic effect on improving the salt shielding property was obtained by the combination of the expansion material and expansive additive. 3.1.2. chloride ion profile the chloride ion profiles of the different concrete samples are shown in figures reffig2 5. the apparent diffusion coefficients and amounts of superficial 257 shinya ito, takeshi iyoda acta polytechnica ctu proceedings figure 2. chloride ion profile for n. figure 3. chloride ion profile for n+ca2. chloride ions calculated by fick’s diffusion equations based on this chloride ion profiles are given in table 3, along with the maximum total chloride ion content in each concrete and the free/immobilised chloride content ratios. the amounts of surface chloride ions estimated from the salinity profiles by fick’s diffusion equations were higher for the n+ca2, n+ca2+ex, and bb samples compared with the n sample as a reference. for the distribution of the chloride ion amount in the depth direction from the penetration surface, the difference in the chloride ion amounts between the surface layer and inside the cured body was large, except for the n sample, and the chloride ion amount decreased with increasing depth. the tendency was remarkable for n+ca2+ex and bb in which ca2 and the expansive additive were used in combination. the apparent diffusivities of the n+ca2+ex, n+ca2, and bb samples were smaller than that of the n sample, by about 42% for the n+ca2 sample, about 54% for the n+ca2+ex sample, and about 69% for the bb sample. this confirmed that chloride ion diffusion was suppressed by using the combination of ca2 and the expansive additive compared with using ca2 alone. in this study, salt water immersion was performed for 48 weeks. therefore, it is considered that the first layer (0 to 10 mm) of the permeation surface, which exhibited figure 4. chloride ion profile for n + ca2 + ex. figure 5. chloride ion profile for bb. the maximum total salt content, was saturated with chloride ions present in the concrete. for the n sample, because the ratios of free/immobilised chloride ions in the depth direction from the penetration surface direction were not significantly different between the surface layer and inside the sample, it is considered that the permeated chloride ions moved and were immobilised at a constant ratio. in contrast, for the n+ca2 and n+ca2+ex samples, the immobilisation ratios of the surface layer and inside the sample were different, and it was inferred that immobilisation and adsorption preferentially occurred. in addition, for the bb sample, although the amount of immobilisation was small, penetration of chloride ions into the surface layer was greatly suppressed, suggesting that the pore structure was such that ion penetration was difficult. therefore, it was considered that the permeation behaviour of chloride ions can be expressed by fick’s diffusion equations for the n sample in which fixation/adsorption and ion permeation proceed at relatively constant rates. however, it is suggested that it is difficult to appropriately describe the permeation behaviour of ions by fick’s diffusion equations for a material that possesses the fixing function, such as ca2, or a material with pore structures that influence ion permeation, such as blast furnace cement. 258 vol. 33/2022 chloride ion penetration behaviour in ca-al concrete fick’s diffusion equation ratio (%) apparent diffusion coefficient (m2/s) chloride ion content at the concrete surface (kg/m3) total chloride ion content (kg/m3) free-chloride ion immobilized chloride ion n 1.02e-11 16.4 14.3 69% 31% n + ca2 5.88e-12 19.8 15.9 57% 43% n + ca2 + ex 4.72e-12 20.9 16.1 60% 40% bb 3.13e-12 22.1 15.0 79% 21% table 3. concrete mix proportion. w/b (%) n 55 opc : 100% n + ca2 opc : 93.5% + ca2 : 6.5% n + ca2 + ex opc : 86.4% + ca2 : 6.5% + ex : 7.1% bb bb : 100% table 4. mix proportions of the pastes. figure 6. hydrate before immersion. 3.2. material characteristics and penetration behaviour of chloride ions the results suggested that the penetration behaviour of chloride ions inside the hardened bodies and the distribution of free chloride ions differed in the concrete formulations, and that the values calculated with fick’s diffusion equations may not necessarily reproduce the actual infiltration behaviour of chloride ions, except for the n sample. therefore, we focused on two material properties: the hydrate and the pore structure. 3.2.1. hydrate the immobilisation performance of chloride ions in each sample was determined. the hydrate was verified by powder x-ray diffraction (xrd) using the paste test specimens with the mix proportions given in table 4. curing of the paste test specimens was performed for 7 days in water. the results are shown in figure 6. compared with the case of no mixing, the intensity the ch peak (10◦ − 11.5◦) of the cement figure 7. hydrate before immersion. hydrate decreased when ca2 or ca2 and the expansive additive were mixed with ordinary cement. this means that ch was consumed by the hydration reactions of ca2 and afm(c) of hc and mono-carbonate or hemi-carbonate formed. the reason why hc and afm(c) were mixed in the reaction of ca2 can be attributed to the small amount of mixed components in the cement. in addition, the same hydrate formed when the combination of ca2 and the expansive additive was used. no peak was observed for ettringite, which is the hydrate of the expansive additive. it was presumed that this was because the so3/al2o3 molar ratio was reduced by the combined use with ca2. to confirm the hydrate change after salt immersion, the paste test specimens with the same mix proportions were immersed in 10% aqueous nacl solution for 2 months. the results of xrd measurements after immersion are shown in figure 7. the peaks of hc and afm, which were observed before salt immersion, disappeared after salt immersion, and they changed to friedel’s salts regardless of the mix proportion. therefore, hc and afm have the function of immo259 shinya ito, takeshi iyoda acta polytechnica ctu proceedings figure 8. friedel’s salt content. figure 9. pore volumes of the four samples. bilising hc and afm as friedel’s salts when chloride ions are present, and it is considered that afm also has immobilising ability even when it is afm(c), such as mono-carbonate and hemi-carbonate. thermal analysis was also performed using paste specimens with the same mix proportions to quantify the amount of friedel’s salt after salt immersion. the amount of friedel’s salt was determined by taking the mass reduction from 247 to 387 ◦c obtained by thermal analysis using the pure synthesised friedel’s salt as 100% and comparing it with the mass reduction rate of the paste in the same temperature region. the amount of friedel’s salt for each sample determined by thermal analysis is shown in figure 8. from the results, compared with the n sample, the amount of friedel’s salt was 1.5 times higher for the n+ca2 sample, 1.45 times higher for the n+ca2+ex sample, and 1.2 times higher for the bb sample, and the immobilisation capacity of chloride ions was remarkably increased in the ca2 blended system. in addition, the expansive additive did not significantly influence the immobilisation ability of chloride ions. 3.2.2. pore structure the influence of the pore structure was investigated, including the migration path of ions, total pore amount, continuous pore amount, and pore network. figure 10. steady-state electrophoresis test results. first, the ratio of the total pore amount to the continuous pore amount measured by the mercury intrusion method using the paste test specimens after salt water immersion was investigated (figure 9). the continuous pore amount was determined by press-fitting mercury twice up to a maximum pressure of 410 mpa, and the mercury press-fit amount at the time of re-pressurisation was the continuous pore amount [4, 5]. from the results, the continuous pore volume of n+ca2+ex was the smallest prior to salt immersion and that of n+ca2 was the smallest after salt immersion. from these results, it is considered that the pores tended to be slightly reduced by using ca2. the absolute value of the total pore amount was in the range 0.3 − 0.4 ml/ml before salt immersion and 0.2 − 0.3 ml/ml after salt immersion. in addition, there was a tendency for the continuous pore volume to slightly decrease after salt immersions for the composites containing ca2 and bb. it was presumed that the effect of densification caused by the change in the hydrate, such as formation of friedel’s salt, affected the continuous pore volume. however, the absolute amount ranged from about 0.09 to 0.15 ml/ml, and it was presumed that the difference was not sufficiently large difference for permeation of chloride ions to greatly change. next, the pore network was investigated. the steady-state electrophoresis test is a technique for evaluating pore networks. it is specified in the "method for testing the effective diffusion factor of chloride ions in concrete by electrophoresis (jsceg-571)" standard of the japan society of civil engineers. it determines the effective diffusion factor from the amount of ions eluted to the anode side by applying a voltage of 15 v to the hardened concrete to accelerate penetration of chloride ions. the results of the steady-state electrophoresis tests for the concrete samples given in table 2 are shown in figure 10. the amount of chloride ions eluted to the anode was in the order n > n + ca2 > n + ca2 + ex > bb, which is the same as the salt immersion test results described in the previous paragraph. however, when the respective behaviours were compared, the chloride ion permeation behaviours were clearly different between 260 vol. 33/2022 chloride ion penetration behaviour in ca-al concrete fixation ability total pore volume continuous pore volume pore network ratios chloride ion suppression effect n 100% 0.26 ml/ml 0.15 ml/ml − − n + ca2 150% 0.20 ∼ 0.26 ml/ml © 0.09 ∼ 0.15 ml/ml © −17% © n + ca2 + ex 145% −25% bb 120% © −52% table 5. summary of material characteristics and penetration behaviour of chloride ions. the sample containing ca2 and the bb sample. here, we will consider the pore structure inside the cured body on the basis of the results obtained from the steady-state electrophoresis test. in the steady-state electrophoresis test, chloride ions are forced to move to the anode side by a potential gradient, and the eluted amount of chloride ions is measured. that is, the ions eluted on the anode side pass through the continuous pores in the cured body, and the eluted amount in the same time depends on the pore network, which is the complexity of the continuous pores. in addition, because the permeation and migration behaviour of chloride ions clearly differed between the immersion method and electrophoresis method for the sample containing ca2, it is considered that the immobilisation function of chloride ions by ca2 is not reflected by the steady electrophoresis method. here, from the relationship between the energisation time and the chloride ion concentration on the anode side shown in figure 10, the cumulative amount of ions since the elution amount of ions became a steady state was approximated by the least squares method. then, assuming that the slope of the approximate straight line is an index representing the pore network, expressed in proportion to the n sample. the elution rate of chloride ions to the anode side were in the order n > n + ca2 > n + ca2 + ex >>> bb. in addition, based on the n sample, the pore network ratios were −52% for the bb sample, −17% for the n+ca2 sample, and −25% for the n+ca2 +ex sample. that is, the pore structures tended to be more complicated using the expansive additive and ca2 than using ca2 alone. therefore, a synergistic effect can be obtained by using the expansive additive and ca2 from the viewpoint of suppressing permeation and migration of chloride ions. 3.2.3. summary of material characteristics and penetration behaviour of chloride ions the summary of study on the effect of material characteristics and penetration behaviour of chloride ions are shown in table 5. 4. summary the material characteristics of concrete containing a combination of ca2 and an expansive additive were investigated as a basic study for durability examination, and the effect of chloride ions on the permeation behaviour was investigated. the following results were obtained. 1. by dividing the amount of chloride ions into immobilised and free chloride ions based on the chloride ion profile obtained by the salt immersion test, the proportions greatly differed depending on the material characteristics. 2. the calculated and measured amounts of surface chloride ions calculated by fick’s diffusion equations differed, except for ordinary concrete. for a material that possesses the immobilisation function, such as ca2, or a material that influences the pore structures, such as blast furnace cement, it is difficult to appropriately describe the penetration behaviour of chloride ions by fick’s diffusion equations. 3. in the evaluation test using paste samples, the same hydrate was produced using ca2 alone and using the combination of the expansive additive and ca2, and friedel’s salt production when chloride ions were present was also equivalent. the combined use of ca2 and the expansive additive had no influence on the immobilisation ability of the chloride ions. 4. the total and continuous pore volumes of each paste specimen were determined by the mercury intrusion method, and there was no significant difference in the permeation of chloride ions. 5. for the pore network calculated on the basis of the steady-state electrophoresis test, the reduction rate of the pore network was larger when the combination of ca2 and the expansive additive was used compared with ca2 alone. it was inferred that the pore structure changed by introduction of expansion strain. 6. although the reduction rate of the pore network using ca2 was smaller than that of the bb sample, the higher immobilisation rate suggested that the same degree of resistance to salt damage was exhibited, albeit with a small total addition rate. references [1] k. tabara, k. yamamoto, m. ashida, et al. capacity of cemented cement hardened with cao and 2al2o3 261 shinya ito, takeshi iyoda acta polytechnica ctu proceedings to fix chloride ions, cement science and concrete technology 64(1):428-34, 2010. https://doi.org/10.14250/cement.64.428. [2] morioka, w. tawara, k. yamamamoto, et al. diffusion inhibiting effect of chloride ion of cao.2al2o3 and its mechanism, concrete technology series no.89 of civil engineering society, report no.2 of the study subcommittee on physical property change and performance evaluation of concrete using admixture materials (333 committee) report no.2, pp.443-448 (2010). [3] k. tsuji, k. suhara. performance evaluation of expanded concrete, technical journal publishing (2011). [4] ryoshi yoshida, kishiji. separation and extraction of continuous pores and ink bottle pores by stepwise injection of mercury, discoloration of samples by mercury injection, summary of lectures at the annual academic lecture of the japan society of civil engineers, 62:5043, 2007. [5] r. yoshida, r. kishi. study on differences in mercury injection processes in cement pastes with different water cement ratios and curing, report on annual articles of concrete engineering, 29:729-734, 2007. 262 https://doi.org/10.14250/cement.64.428 doi:10.14311/app.2021.30.0087 acta polytechnica ctu proceedings 30:87–92, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app beam element under finite rotations emma la malfa ribolla∗, milan jirásek, martin horák czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: emma.la.malfa.ribolla@fsv.cvut.cz abstract. the present work focuses on the 2-d formulation of a nonlinear beam model for slender structures that can exhibit large rotations of the cross sections while remaining in the small-strain regime. bernoulli-euler hypothesis that plane sections remain plane and perpendicular to the deformed beam centerline is combined with a linear elastic stress-strain law. the formulation is based on the integrated form of equilibrium equations and leads to a set of three first-order differential equations for the displacements and rotation, which are numerically integrated using a special version of the shooting method. the element has been implemented into an open-source finite element code to ease computations involving more complex structures. numerical examples show a favorable comparison with standard beam elements formulated in the finite-strain framework and with analytical solutions. keywords: finite rotations, nonlinear beam, shooting method. 1. introduction highly slender fiberor rod-like components represent essential constituents of mechanical systems in many fields of application such as civil, mechanical and biomedical engineering. it is widely recognized that slender bodies can be efficiently modeled applying a beam theory instead of a 3d continuum mechanics theory. kirchoff proposed the first beam formulation which includes large three-dimensional deformations [1, 2], and reissner completed the theory for both two-dimensional [3] and three-dimensional cases [4] with two additional deformation measures representing the shear distortion of beam segments. reissner’s finite-strain beam theory is one of the most important geometrically nonlinear models, subsequently extended and used by many other authors for 2-d and 3-d analysis of static as well as dynamic problems. simo developed a dynamic formulation for reissner’s beam [5] and together with vu-quoc [6] initiated the finite element implementation. he also introduced the useful concept of a geometrically exact beam, based on recasting reissner’s theory in a form which is valid for any magnitude of displacements and rotations. in this paper, a geometrically nonlinear beam model is formulated for the 2-d case. this model is applicable when the strains remain in a limited range while the rotations of beam sections can become arbitrarily large, and it properly accounts for the effect of curvature on the change of distance between end sections measured along the chord. therefore the model considers large rotations and displacements while strains are treated as small and simple hooke’s law is used (an extension to a nonlinear material law would be relatively straightforward). cross sections are still assumed to remain planar and perpendicular to the deformed beam axis. the formulation exploits the equilibrium equations in their strong form. they are combined with the kinematic equations and generalized material equations (linking internal forces to the curvature and centerline extension) and then numerically integrated. 2. model description 2.1. basic kinematic assumptions in this section, the kinematics of the element is presented, leading to the strain-displacement equations. each beam is analyzed in its local coordinate system, with the x-axis passing through the end joints in the undeformed configuration and the z-axis rotated by 90◦ clockwise, while the rotations are positive counterclockwise (see figure 1). we use u and w for horizontal and vertical displacements at an arbitrary point, while the displacements at the centerline are denoted by us and ws. let us consider an infinitesimal segment of the centerline of initial length dx, which is after deformation transformed into a segment of length dx̄ =! (dx + dus)2 + dw2s . the corresponding centerline stretch can be expressed as λs = dx̄ dx = ! (1 + u′s)2 + w′2s (1) where the prime denotes differentiation with respect to x. the assumption of planarity of the deformed section and of its perpendicularity to the deformed centerline implies that the rotation angle is linked to the centerline displacements by the relation tan ϕ = − dws dx + dus = − w′s 1 + u′s (2) 87 http://dx.doi.org/10.14311/app.2021.30.0087 http://ojs.cvut.cz/ojs/index.php/app e. la malfa ribolla, m. jirásek, m. horák acta polytechnica ctu proceedings figure 1. kinematics of deformation of the presented nonlinear beam model. and that the displacements of a general point with coordinate z can be expressed as u = us + z sin ϕ (3) w = ws − z(1 − cos ϕ) (4) in contrast to the geometrically linear beam model with u = us + zϕ and w = ws, here the vertical displacement varies along the height of the section. however, the change of distance from the centerline due to deformation perpendicular to the beam axis is still neglected. differentiating (3)–(4) with respect to x, we obtain u′ = u′s + z cos ϕ · ϕ ′ (5) w′ = w′s − z sin ϕ · ϕ ′ (6) and we can express the longitudinal stretch at an arbitrary location λ = ! (1 + u′)2 + w′2 = = ! λ2s + 2zλsϕ′ + z2ϕ′2 = λs + zϕ ′ (7) it is worth noting that the stretch is a linear function of the spatial coordinate z. consequently, if we use a linear stress-strain law formulated in terms of the biot strain defined as ε = λ − 1, the resulting work-conjugate stress will be a linear function of z, too. the choice of working with linear stress-strain relation between the biot strain and stress can be justified in a regime where the strains are assumed to remain small. making use of (7), we can express the biot strain as ε = λ − 1 = λs − 1 + zϕ′ = εs + zκ (8) where εs = λs − 1 = ! (1 + u′s)2 + w′2s − 1 (9) is the strain at the centerline and κ = ϕ′ (10) is the curvature. 2.2. stress resultants and equilibrium based on (8) combined with hooke’s law σ = e ε, quantities εs and κ that characterize the deformation of an infinitesimal beam segment can be linked to the normal force n and bending moment m by the standard linear relations n = " a σ da = " a e(εs + zκ) da = ea εs (11) m = " a zσ da = " a ez(εs + zκ) da = ei κ (12) in which ea is the sectional axial stiffness and ei is the sectional bending stiffness. as usual, e denotes young’s modulus, a is the sectional area and i is the sectional moment of inertia. the structure is assumed to be loaded only at its joints, i.e., no external loads are applied on individual beams, and so the internal forces in each beam can be expressed in terms of the end forces and moments, based on equilibrium conditions written for the deformed configuration. from equilibrium of the finite beam segment between the left end section and a generic section x, one can, according to fig. 2, derive expressions n (x) = −xab cos ϕ(x) + zab sin ϕ(x) (13) m (x) = −mab + xabws(x) − zab(x + us(x)) (14) in which xab, and zab are the left-end forces (positive to the right and downwards) and mab is the left-end moment (positive counterclockwise). for the sake of brevity, we will consider the moment as one of the (generalized) forces, and, in a similar spirit, the rotation as one of the (generalized) displacements. since the internal forces are involved only algebraically, they can be easily eliminated and equations (2) and (9)–(14) can be converted into the following set of first-order differential equations for unknown functions us, ws and ϕ: ϕ′(x) = −mab + xabws(x) − zab(x + us(x)) ei (15) u′s(x) = −1 + cos ϕ(x)+ + −xab cos ϕ(x) + zab sin ϕ(x) ea cos ϕ(x) (16) 88 vol. 30/2021 beam element under finite rotations m(x) n(x) x z xab zab mab x us ws ! x̃ z̃ ã a figure 2. schematic representation of beam internal forces at a general cross section x and left-end forces and moment. w′s(x) = − sin ϕ(x)+ + xab cos ϕ(x) − zab sin ϕ(x) ea sin ϕ(x) (17) these equations are accurate even for very large rotations. the only limitation is that the strain remains small, which is true if εs ≪ 1 and hκ ≪ 1 where h is the depth of the cross section. in the context of a finite element analysis, the joint displacements are global degrees of freedom that are determined by equilibrium iterations on the structural level. in each iteration, one needs to determine, for each beam, the end forces that correspond to the given joint displacements and rotations. this means that constants xab, zab and mab in (15)–(17) are not known in advance. on the other hand, the displacements at both end sections are prescribed. at the left end, conditions ϕ(0) = ϕa (18) us(0) = ua (19) ws(0) = wa (20) can be considered as initial conditions for first-order differential equations (15)–(17). if the values of xab, zab and mab are somehow estimated, these equations can be integrated, at least numerically, and the values of ϕ(l), us(l) and ws(l) can be determined. the values of forces xab, zab and mab then need to be adjusted such that the resulting kinematic quantities at the right end satisfy the remaining boundary conditions ϕ(l) = ϕb (21) us(l) = ub (22) ws(l) = wb (23) for a given set of end displacements, the initial estimate of the left-end forces can be constructed based on linear beam theory, or on the values at the end of the previous step if the calculation is done in the context of an incremental iterative structural analysis. in fact, the suggested approach is a special version of the shooting method. normally, the shooting method iterates on some of the initial values. actually, the left-end forces would appear in static boundary conditions at the left end, which make it possible to integrate the differential equations of equilibrium and proceed from external forces to internal forces. here, the equilibrium equations were directly presented in their integrated form (13)–(14), and the unknown variables play the role of integration constants instead of boundary values. 2.3. numerical solution the suggested numerical approach is based on the full form of equations (15)–(17). the right-end displacements ub, wb and ϕb can be evaluated if the left-end displacements ua, wa and ϕa and the left-end forces xab, zab and mab are given. to this end, the derivatives in (15)–(17) can be replaced by finite differences and the interval [0, l] is divided into n equally sized numerical segments. mapping of the right-end displacements, rotation, forces and moment based on the left-end displacements and rotation can formally be described by ub = g(f ab, ua) (24) where ua = # $ ua wa ϕa % & , ub = # $ ub wb ϕb % & , (25) f ab = # $ xab zab mab % & and g is a column matrix of three functions of six variables. for given values of ua and ub, equation (24) represents a set of three nonlinear algebraic equations for unknowns f ab. the solution is found by the newton-raphson method, using the recursive formula δf (k) ab = g −1 ' f (k) ab , ua ( ' ub − g ' f (k) ab , ua (( (26) f (k+1) ab = f (k) ab + δf (k) ab , k = 0, 1, 2, . . . (27) where g = ∂g ∂f ab (28) is the jacobi matrix of mapping g with respect to variables f ab. the entries of the jacobi matrix are evaluated numerically using the linearized version of the computational scheme. once the left-end forces have converged, the right-end forces are easily evaluated from equilibrium conditions written for the whole beam. the foregoing equations were formulated in a local coordinate system aligned with the undeformed 89 e. la malfa ribolla, m. jirásek, m. horák acta polytechnica ctu proceedings m figure 3. pure bending of a cantilever beam subjected to end moment. beam. since the deformed shape of the beam and the distribution of internal forces must be invariant with respect to rigid-body translations and rotations, one can also use a rotated coordinate system x̃ − z̃ with the origin located at the left end of the beam in the deformed configuration (point ã in fig. 2) and with the x̃ axis in the direction of the tangent to the deformed centerline at the left end. with respect to this coordinate system, the left-end displacements are always zero, and so ua can be omitted from the list of arguments of function g. of course, the rightend displacements ub must be computed from the global components of the joint displacements using an appropriate transformation rule, and the end forces obtained by the iterative process (26)–(27) must be transformed back into the global coordinate system. consistent linearization of the relation between the end forces and end displacements leads to the element tangent stiffness matrix. details will be presented in a full journal paper. 3. numerical examples a nonlinear beam element based on the proposed approach has been implemented into oofem [7, 8], an object-oriented finite element code. to verify the implementation and demonstrate the potential of the suggested approach, two examples involving simple structures are shown next. 3.1. pure bending of a cantilever beam the first test is a cantilever of length l = 8 and bending stiffness ei = 1 loaded by a concentrated end moment m on its right-end. the exact solution to this problem is a circular arc with radius r = ei/m . an applied end moment, m = π/4, will force the rod to deform into a full closed circle. in this example, the moment is applied in six load steps, making the rod wind around itself at the end of the sixth step. the deformed shape of the beam at the end of each step is depicted in fig. 3. the solution of the presented figure 4. close-up view of the solution at the end of the sixth step using 6,10,20, and 50 segments. nonlinear model is compared with the one obtained by employing the geometrically exact finite beam element by simo and vu-quoc [6] with a mesh of eight elements. the exact solution is reported as well. the overall agreement is good, and the simulation based on the present model, which uses only one two-noded element, is closer to the analytical solution. the example demonstrates that the present model allows for a dramatic reduction of the number of global degrees of freedom, but of course the number of segments n used for numerical integration of the governing equations (15)–(17) must be chosen high enough to provide a good approximation. the presented results have been obtained using 100 segments. a close-up view of the sixth step circle is showed in fig. 5 for calculations in which 6, 10, 20 and 50 numerical segments were employed. to ease the interpretation of the results we plot straight segments in between the integration points even though the curvature is constant along the beam. it turns out that 20 segments provide an acceptable solution, which is further improved and gets close to the analytical one if 50 segments are used. it is also noted that even with just 10 segments the method provides an accurate estimation of displacements at the integration points. to better quantify the previous comparisons, we calculated the sum of the distances between the the analytical and the computed solutions, normalized with respect to the number of sample points. the results for the sixth load step are reported in table 1. already for 8 integration segments, the error is smaller that 5%. the results indicate that comparable accuracy can be obtained if the present method replaces traditional finite elements by the same number of integration segments located within one single finite element. 3.2. clamped-hinged circular arch subject to point load the second example deals with an arch instability after large asymmetrical deflections. this specific 90 vol. 30/2021 beam element under finite rotations model error [%] simo and vu-quoc [6] 2.98 present approach, 101 segments 0.23 present approach, 50 segments 0.29 present approach, 20 segments 0.70 present approach, 10 segments 2.13 present approach, 8 segments 3.17 present approach, 6 segments 5.39 table 1. sum of the absolute values of the differences between the exact and the computed displacement solutions calculated for the sixth load step. the model of simo and vu-quoc [6] uses an eight-element mesh while the presented beam model uses one element with a variable number of segments of the integration scheme. figure 5. circular arch geometry. problem was investigated by many authors [6, 9] and the exact solution based on the kirchhoff-love theory was given by dadeppo and schmidt [10]. the arch is circular with one boundary clamped and the other boundary hinged, and is loaded by a point load, as shown in fig. 5. the cross section of the beam is a rectangle of depth h = 2.289 and width t = 1.0. the elastic modulus is set to e = 1.0 × 106 and the poisson ratio is zero. geometrical and constitutive parameters are dimensionless meaning that any length or force scale can be used for these values. simo and vu-quoc [6] performed a fe analysis with a forty-element mesh, leading to a buckling load of 9.0528 ei/r2, while wood and zienkiewicz [9] calculated a buckling load of 9.24 ei/r2. the exact value of the buckling load reported by dadeppo and schmidt [10] is 8.97 ei/r2. in our simulations, we used forty elements as well, each with four integration segments, and we obtained a buckling load of 8.9874 ei/r2, which is very close to the analytical solution and exhibits a relative error less than 0.02%. the full comparison with the analytical solution is reported in fig. 6 where we plot the load-displacement and the load-rotation curves relative to the top of the arch (loaded point). again, the analytical solution is satisfactorily reproduced. the deformed configurations for the four load levels numbered in the curves are also shown together with the bending moments indicated by colors. the calculation was completed in 612 load steps using the arc-length method with a 1 2 3 4 1 2 3 4 1 2 3 1 2 3 4 figure 6. load-displacements and load-rotation of the top of the arch (top). grey-dashed line represents the analytical solution reported by dadeppo and schmidt [10]. deformed shapes and moment trend along the arch relative to the four load levels numbered in the top curves. maximum number of 3 iterations per step. it is noted that simo and vu-quoc [6] and wood and zienkiewicz [9] show a longer snap-through behavior together with the associated deformed shapes even though the response is not reasonable because of non-physical penetration of the left support into the structure. this aspect was investigated by simo et al. [11], who showed that a contact constraint condition on the left support needs to be introduced to obtain a more realistic solution. for the aim of 91 e. la malfa ribolla, m. jirásek, m. horák acta polytechnica ctu proceedings this paper we did not consider contact activation and the subsequent stiffening effect in the structure even though this extension could be added. 4. conclusions we present an efficient formulation for a geometrically nonlinear beam model that accounts for arbitrarily large rotations of the beam sections and the effect of curvature on the change of distance between end sections measured along the chord. in the first stage, the model is restricted to the small-strain framework and cross sections are still assumed to remain planar and perpendicular to the deformed beam axis. equilibrium equations in their strong form are combined with the kinematic equations and generalized material equations and then numerically integrated using a special version of the shooting method. two examples show the capabilities of the element to replicate analytical results in a more accurate way with respect to existing beam finite elements that account for geometric nonlinearities. moreover, the h-refinement of finite elements, which increases the number of global unknowns and is thus computationally demanding, can be substituted by employing a proper number of segments in the integration scheme, while keeping the number of elements limited. the element will be extended into a formulation that takes into account an initial curvature and also accounts for a possible contact mechanism activation. acknowledgements the authors acknowledge the support of the czech science foundation (project no. 19-26143x). references [1] g. kirchhoff. ueber das gleichgewicht und die bewegung eines unendlich dünnen elastischen stabes. journal für die reine und angewandte mathematik 1859(56):285–313, 1859. [2] e. h. dill. kirchhoff’s theory of rods. archive for history of exact sciences 44(1):1–23, 1992. [3] e. reissner. on one-dimensional finite-strain beam theory: the plane problem. zeitschrift für angewandte mathematik und physik zamp 23(5):795–804, 1972. [4] e. reissner. on finite deformations of space-curved beams. zeitschrift für angewandte mathematik und physik zamp 32(6):734–744, 1981. [5] j. c. simo. a finite strain beam formulation. the three-dimensional dynamic problem. i. computer methods in applied mechanics and engineering 49(1):55–70, 1985. [6] j. c. simo, l. vu-quoc. a three-dimensional finite-strain rod model. part ii: computational aspects. computer methods in applied mechanics and engineering 58(1):79–116, 1986. [7] b. patzák, z. bittnar. design of object oriented finite element code. advances in engineering software 32(10-11):759–767, 2001. [8] b. patzák. oofem—an object-oriented simulation tool for advanced modeling of materials and structures. acta polytechnica 52(6), 2012. [9] r. d. wood, o. zienkiewicz. geometrically nonlinear finite element analysis of beams, frames, arches and axisymmetric shells. computers & structures 7(6):725–735, 1977. [10] d. dadeppo, r. schmidt. instability of clamped-hinged circular arches subjected to a point load. journal of applied mechanics, transactions asme 42(4):894–896, 1975. [11] j. simo, p. wriggers, k. schweizerhof, r. taylor. finite deformation post-buckling analysis involving inelasticity and contact constraints. international journal for numerical methods in engineering 23(5):779–800, 1986. 92 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0085 acta polytechnica ctu proceedings 34:85–93, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague adhesion test – testing of selected adhesives on fired clay kristýna richterová∗, pavel heinrich, petr bílý czech technical university in prague, faculty of civil engineering, department of concrete and masonry structures, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: kristyna.richterova@fsv.cvut.cz abstract. the following paper deals with the comparison of adhesion of specimens made of fired clay and selected adhesives. the adhesion was tested on specimens at the age of one, two and three days in order to determine the parameters of young masonry. nine different adhesives (adhesives intended for masonry and adhesives for others material) were tested in total. the above mentioned test is not standardized. the obtained results from the adhesion test determine the tensile strength between the adhesive and the fired clay and the adhesion of selected adhesive to the fired clay. the small specimens were made of two ceramic plates and an adhesive. selected adhesives were applied to ceramic plates according to the usual standards. eighty-one small specimens were produced in total – twenty-seven specimens for the adhesion test at each defined age of the specimens (one day, two days, three days). the produced specimens were cured in the climatic chamber with set laboratory conditions until the time of the test. the specimens were tested in the laboratory of the heluz brick plant in dolní bukovsko. this paper deals with the procedure of preparation of the specimens, the course of the test and the comparison of gained results. keywords: fired clay, adhesion test, adhesive. 1. introduction the adhesion test determines the tensile strength between the adhesive and the building material (in our case fired clay) and the adhesion of selected adhesive to the building material (fired clay). although requirement for tensile strength between a fired clay and adhesive is not defined by any regulation, the obtained results of the measurement provide information, which is important for the design of innovative construction methods. the test was performed at the age of the samples of one, two and three days. the measured values of “young” samples give an idea of the behaviour of selected adhesives available on the czech market and indicate, which adhesives reach the required strength in a short time. a similar adhesion test (the centric adhesion test on two-unit test samples with 2c-pur adhesive and thin-layer mortar) was performed in 2013 by the institute of building materials research (ibac) in aachen. the performed tests were a necessary basis for the introduction of a new construction method in germany – see [1]. today´s trend is the construction of structures in the shortest possible time, using suitable and efficient construction methods in compliance with high quality implementation, so the adhesion test is a felicitous tool for determining the required properties of tensile stressed structures. 2. materials and samples eighty-one small samples were produced for the adhesion test – twenty-seven samples for adhesion test at each defined age of the samples. the samples were made of fired clay and selected adhesive. nine adhesives were tested in total. the selected adhesives are listed in the following table. sample description a heluz sidi b heluz sb mortar for a thin joint c heluz tytan pur foam (thin-filmadhesive) e pu-contact two-component adhesive f den braven – multi kleber –masonry foam g den braven – low expansion – masonryfoam i den braven – mamut glue (hightack) j illbruck – universal chemical anchorot120 k den braven – wood glue – woodfixd3 table 1. list of used adhesives [2]. the fired clay in the form of plates used for the adhesion test – size 30×30 mm and thickness 15 mm – has the same properties as heluz bricks. small samples for the adhesion test were made of two plates of fired clay and selected adhesive – table 1. heluz sidi is a silicate-dispersion masonry mortar for a thin joint, which is specified for masonry made of clay blocks heluz grinded. heluz sidi was 85 https://doi.org/10.14311/app.2022.34.0085 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en k. richterová, p. heinrich, p. bílý acta polytechnica ctu proceedings launched in the spring of 2020, it is a prepared mixture and is applied evenly crosswise with a structured roller [3]. heluz sb mortar for a thin joint is applied to the ribs of the grinded clay blocks and is applied with heluz applicator roller sb or by direct dipping of the bed joint in a pre-prepared mixture. the mixture is prepared from heluz sb and water – in a ratio of 0.45 liters of water per 1 kg heluz sb. heluz sb was applied to small samples on the whole surface of the bed joint using a spatula [4]. heluz tytan pur foam was applied with the application gun on the whole surface of the bed joint [5]. pu-contact two-component adhesive is composed of two components – adhesive and accelerator – and it was applied on the whole surface of the bed joint of both parts of the sample. the adhesive was applied by a pneumatic spray gun. the application of pucontact two-component adhesive has to be fast due to the presence of the accelerator. the increase of adhesion and application of the adhesive are very fast. den braven masonry foams were selected to compare the results with heluz tytan pur foam and were also applied by application gun [6, 7]. all samples with masonry foam were not loaded fast enough – the foam has swelled up, and the thickness of bed joint was greater than stated in the technical datasheets of adhesives. swelling of the foam negatively affects its declared values, the probability of failure at the site of adhesive increases and therefore the values obtained from the adhesion test are on the side of safety. mamut glue is a one-component adhesive ms polymer-based. this adhesive has a declared high initial adhesion (up to 500 kg/m2) and has been applied by application gun [8]. illbruck – universal chemical anchor ot120 is a two-component resin without styrene. this chemical anchor is intended for anchoring rods, bolts and bushings to concrete and common building materials. this chemical anchor is also suitable for anchoring in hollow masonry. the chemical anchor is applied by an application gun for silicone sealants [9]. the wood glue – woodfix d3 – is an aqueous dispersion glue and is intended for gluing wood [10]. figure 1. the fired clay in the form of plates and the produced samples. the produced samples (9×3×3) for the adhesion test were cured in the climatic chamber until the time of the test. the climatic chamber was set to figure 2. the selected adhesives: b – heluz sb mortar for a thin joint, a – heluz sidi, c – heluz tytan pur foam (thin-film adhesive), e – pu-contact two-component adhesive. figure 3. the selected adhesives: f – multi kleber – masonry foam, g – low expansion – masonry foam, i – mamut glue (high tack), j – illbruck – universal chemical anchor ot120, k – wood glue – woodfix d3. laboratory conditions – constant temperature 23 ◦c, constant relative humidity 50 %. the test samples were loaded by steel plate (m = 280 g) in the chamber. the steel plate is a component for the adhesion test. each sample from a series of test samples was equipped with steel plates with hitches for attachment to the test apparatus before the adhesion test. steel plates measuring 50×50 mm were glued by epoxy glue (sikadur®-31 cf rapid) to the lower and upper surface of the test sample. a small number of samples with the selected adhesive at each defined age (three test samples) were selected due to the total number of test samples (eightyone test samples).the adhesion test was performed mainly for the selection of unsuitable adhesives. adhesives with satisfactory results from this test will be subjected to further testing in the future. 86 vol. 34/2022 adhesion test – testing of selected adhesives on fired clay figure 4. application of epoxy to samples for fixing steel plates. 3. adhesion measurement the adhesion test was performed at the age of the samples one day, two and three days – each series of 27 samples at a defined age. the series of samples at a defined age was transferred from the climatic chamber to the testing laboratory in dolní bukovsko on the day of the test. at first the cross-sectional area ai of each test sample was measured using caliper. then the samples were placed in a test apparatus and after that loaded with a tensile force. the used test apparatus was apparatus named tiratest 2803, which can exert a force of up to 3 kn. this force is sufficient for the purpose of the adhesion test. figure 5. the sample preparation in a test apparatus. at first the sample in a test apparatus was fixed manually. after that the sample was loaded with tensile force – the value of increase of tensile force was set to 10 n/s – until the sample failed. the measured tensile force fti,max and the type of failure of the test sample were recorded in the prepared measurement table. the possible types of failure of the sample are failure in the adhesive, in the fired clay, in a combination of fired clay and adhesive or in the epoxy between the fired clay and the steel plate. the named failure types and scheme of adhesion test are depicted in the following figures. from the tensile force fti,max and from the crosssectional area ai the tensile stress σti, the average stress σtm and finally the characteristic stress σtk were calculated. the characteristic stress σtk was figure 6. the scheme of adhesion test. figure 7. the possible types of failure of the sample in the adhesion test. calculated using the standard deviation s, the coefficient of variation v and the uncertainty coefficient kn (2.01 (n=2), resp. 1.89 (n=3)). the individual parameters were calculated according to the following formulas. ai = a · b [mm2] (1) σti = fti,max a [mpa] (2) σtm = 1 n · ∑ σti [mpa] (3) σtk = σtm · (1 − kn · v ) (4) s = √ 1 n − 1 · ∑ (σti − σtm)2 (5) v = s σtm (6) 87 k. richterová, p. heinrich, p. bílý acta polytechnica ctu proceedings 4. results and discussion the values and the information obtained from the adhesion test in the following tables state which adhesives performed the best in the adhesion test. the assessment of the adhesion test depends mainly on the measured values of the tensile force at the failure of the sample, but the variance of the results and the character of the failure of the sample are also important. the type of sample failure gives us an idea of what limits the measured value of tensile stress and whether the used adhesive can resist greater tensile stress then was measured. the failure of the sample in the fired clay says that the used adhesive is likely to resist greater tensile stress than the fired clay used to production of the samples; however, this type of failure did not prevail in any set of samples. the samples that demonstrated satisfactory results in the adhesion test are the samples made of adhesives den braven – mamut glue (high tack), heluz sidi, pu-contact two-component adhesive and den braven – wood glue – woodfix d3. mamut glue (high tack) performed the best in the adhesion test and the samples with this adhesive failed in the adhesive or in a combination of fired clay and the adhesive – it can be assumed that the adhesive is limited by tensile stress values close to the values measured by us. the negative factors of mamut glue (high tack) are the unpredictability of sample failure and its price. a typical failure of pu-contact two-component adhesive is failure in the adhesive at different deformation. this adhesive is very malleable, with reduced adhesion when applied unprofessionally and is applied in a thin layer. very surprising results were measured on samples with applied wood glue – woodfix d3. the measured values of tensille stress were higher for wood glue than for some samples with aplicated adhesives intended for masonry. the low expansion foam (den braven) was the worst of the tested masonry foams. the worst of all the samples were those with the adhesive illbruck – universal chemical anchor ot120 and heluz sb mortar for a thin joint. the principle of the chemical anchor consists in its expansion into the borehole; this was not possible in the case of application to the contact area of small samples. the samples with the chemical anchor were broken in the adhesive in all cases, which did not adhere sufficiently and reached low resistance to tensile stress. the performed adhesion test is not completely valid, because the number of test samples for series at a defined age was small (three samples for every adhesive at a defined age). relatively high standard deviations in individual series of measurements (small number of test samples) caused a significant decrease of the characteristic stress of some samples aged two days and three days, respectively. the large variance of the measured values is also the reason for the negative value of the characteristic stress of samples with adhesive named heluz sb mortar for a thin joint at the age of one day. it would be necessary to produce and test more small samples to obtain valid values of the characteristic stress. 5. conclusion this article introduces the results of the adhesion test performed on the small samples, which were made of fired clay in the form of plates and selected adhesive. the obtained results are not completely valid (a small number of test samples), but the results give an idea of the tensile properties of selected adhesives applied to the fired clay. therefore they are useful for the selection of adhesives for a future, more detailed study. the adhesion test was performed at an early age of the samples to determine the adhesion and tensile strength between the fired clay and the selected adhesive. these ascertained properties are important for the progress of innovative construction methods – not only for the construction of brick buildings. surprising results were measured on samples with applied adhesives that are not intended for masonry. we must be careful about these results because we do not know the development of their properties in contact with masonry over a longer period of time. the research will continue by deeper examination of several selected adhesives. list of symbols a,b horizontal dimension of the sample ai cross-sectional area fti,max tensile force n number of measurements σti tensile stress σtm average stress σtk characteristic stress s standard deviation v coefficient of variation kn uncertainty coefficient acknowledgements the acknowledgement belongs to assoc. prof. petr bílý, who was the supervisor of my master thesis and to the company heluz, which helped to perform the adhesion test. the preparation of this paper was financially supported by student grant of the czech technical university in prague sgs19/149/ohk1/3t/11. 88 vol. 34/2022 adhesion test – testing of selected adhesives on fired clay a – heluz sidi day name crosssectional area tensile force tensile stress type of failure nr. of measurements avg. stress std. deviation char. stress ai [mm2] fti,max [n] σti [mpa] n σtm [mpa] s [mpa] σtk [mpa] 1 a1 961.61 543.22 0.565 adhesive 2 0.583 0.025 0.532a2 959.31 575.86 0.600 adhesive a3 951.04 adhesive(when clamping) 2 a1 956.97 958.41 1.002 3/4 adhesive + 1/4 fired clay 3 0.954 0.171 0.630 a2 945.87 1036.55 1.002 fired clay a3 944.93 721.25 0.763 fired clay 3 a1 945.55 779.50 0.824 fired clay 3 0.687 0.219 0.273a2 965.87 773.15 0.801 fired clay a3 960.64 417.10 0.434 4/5 adhesive + 1/5 fired clay b – heluz sb mortar for a thin joint day name crosssectional area tensile force tensile stress type of failure nr. of measurements avg. stress std. deviation char. stress ai [mm2] fti,max [n] σti [mpa] n σtm [mpa] s [mpa] σtk [mpa] 1 b1 953.53 270.47 0.284 adhesive 2 0.341 0.224 -0.082b2 952.96 144.19 0.151 adhesive b3 945.78 555.59 0.587 adhesive 2 b1 951.72 219.49 0.231 adhesive 3 0.347 0.131 0.099b2 942.94 302.15 0.320 adhesive b3 946.13 462.38 0.489 adhesive 3 b1 946.79 when clamping 2 0.275 0.050 0.175b2 966.27 300.20 0.311 adhesive b3 956.36 229.42 0.240 adhesive c – heluz tytan pur foam (thin-film adhesive) day name crosssectional area tensile force tensile stress type of failure nr. of measurements avg. stress std. deviation char. stress ai [mm2] fti,max [n] σti [mpa] n σtm [mpa] s [mpa] σtk [mpa] 1 c1 956.04 325.55 0.341 adhesive 3 0.338 0.163 0.031c2 956.31 477.45 0.499 adhesive c3 935.68 162.82 0.174 adhesive 2 c1 958.21 177.41 0.185 adhesive 3 0.192 0.006 0.180c2 954.40 183.67 0.192 adhesive c3 952.93 188.39 0.198 adhesive 3 c1 955.73 308.77 0.323 adhesive 3 0.297 0.034 0.233c2 956.85 296.97 0.310 adhesive c3 942.39 243.62 0.259 adhesive 89 k. richterová, p. heinrich, p. bílý acta polytechnica ctu proceedings e – pu-contact two-component adhesive day name crosssectional area tensile force tensile stress type of failure nr. of measurements avg. stress std. deviation char. stress ai [mm2] fti,max [n] σti [mpa] n σtm [mpa] s [mpa] σtk [mpa] 1 e1 944.30 adhesive(when clamping) 2 0.724 0.078 0.568e2 964.10 644.85 0.669 adhesive e3 964.41 750.96 0.779 adhesive 2 e1 939.39 526.22 0.560 adhesive 3 0.443 0.103 0.248e2 947.68 382.27 0.403 3/4 adhesive + 1/4 fired clay e3 926.90 338.60 0.365 adhesive 3 e1 980.25 27.68 0.028 adhesive(did not adhere) 2 0.534 0.038 0.457e2 946.78 530.87 0.561 adhesive e3 946.88 479.58 0.506 adhesive f – den braven – multikleber – mansory foam day name crosssectional area tensile force tensile stress type of failure nr. of measurements avg. stress std. deviation char. stress ai [mm2] fti,max [n] σti [mpa] n σtm [mpa] s [mpa] σtk [mpa] 1 f1 991.60 211.40 0.213 adhesive 3 0.211 0.048 0.121f2 968.45 157.38 0.163 adhesive f3 962.23 248.28 0.258 adhesive 2 f1 723.23 368.26 0.509 adhesive 3 0.393 0.137 0.133f2 730.33 176.21 0.241 adhesive f3 948.33 406.53 0.429 adhesive 3 f1 859.60 253.21 0.295 adhesive 3 0.426 0.167 0.110f2 744.07 457.25 0.615 adhesive f3 948.64 350.27 0.369 adhesive g – den braven – low expansion – masonry foam day name crosssectional area tensile force tensile stress type of failure nr. of measurements avg. stress std. deviation char. stress ai [mm2] fti,max [n] σti [mpa] n σtm [mpa] s [mpa] σtk [mpa] 1 g1 969.70 182.85 0.189 adhesive 3 0.165 0.044 0.082g2 956.29 109.63 0.115 adhesive g3 855.50 164.46 0.192 adhesive 2 g1 885.25 94.46 0.107 adhesive 3 0.198 0.084 0.040g2 989.78 267.68 0.270 adhesive g3 755.55 164.08 0.217 adhesive 3 g1 969.96 328.94 0.339 adhesive 3 0.417 0.190 0.058g2 915.55 255.04 0.279 5/6 adhesive + 1/6 fired clay g3 822.25 521.58 0.634 5/6 adhesive + 1/6 fired clay 90 vol. 34/2022 adhesion test – testing of selected adhesives on fired clay i – den braven – mamut glue (hight tack) day name crosssectional area tensile force tensile stress type of failure nr. of measurements avg. stress std. deviation char. stress ai [mm2] fti,max [n] σti [mpa] n σtm [mpa] s [mpa] σtk [mpa] 1 i1 937.27 1098.71 1.172 adhesive 2 1.042 0.118 0.819i2 979.38 922.21 0.942 adhesive i3 980.92 992.69 1.012 adhesive 2 i1 938.44 1076.41 1.147 3/4 adhesive + 1/4 fired clay 3 1.004 0.176 0.672 i2 980.00 791.50 0.808 2/3 adhesive + 1/3 fired clay i3 965.33 1022.11 1.059 adhesive 3 i1 957.52 560.67 0.586 fired clay 3 0.769 0.159 0.468i2 974.93 841.90 0.864 adhesive i3 958.48 823.55 0.859 1/3 adhesive + 2/3 fired clay j – illbruck – universal chemical anchor ot120 day name crosssectional area tensile force tensile stress type of failure nr. of measurements avg. stress std. deviation char. stress ai [mm2] fti,max [n] σti [mpa] n σtm [mpa] s [mpa] σtk [mpa] 1 j1 959.45 158.83 0.166 adhesive 2 0.244 0.110 0.022j2 974.06 adhesive(when clamping) j3 1005.73 323.34 0.321 adhesive 2 j1 1003.29 214.63 0.214 adhesive 3 0.248 0.029 0.192j2 667.99 176.98 0.265 adhesive j3 970.94 256.95 0.265 adhesive 3 j1 942.78 265.29 0.281 adhesive 2 0.317 0.050 0.217j2 962.20 338.38 0.352 adhesive j3 962.54 adhesive(when clamping) 91 k. richterová, p. heinrich, p. bílý acta polytechnica ctu proceedings k – den braven – wood glue – woodfix d3 day name crosssectional area tensile force tensile stress type of failure nr. of measurements avg. stress std. deviation char. stress ai [mm2] fti,max [n] σti [mpa] n σtm [mpa] s [mpa] σtk [mpa] 1 k1 947.99 344.52 0.363 adhesive 2 0.449 0.121 0.206k2 955.68 2.84 0.003 fired clay(when clamping) k3 981.25 524.34 0.534 adhesive 2 k1 974.64 980.99 1.007 adhesive 3 0.878 0.179 0.540k2 961.84 647.83 0.674 fired clay k3 956.02 912.48 0.954 adhesive 3 k1 981.08 440.06 0.449 adhesive 3 0.659 0.210 0.263k2 992.86 656.44 0.661 fired clay k3 946.76 822.06 0.868 3/4 adhesive + 1/4 fired clay table 2. the values and the information obtained from the adhesion test [1]. figure 8. summary of results from the adhesion test at the age of the samples one, two and three days [1]. 92 vol. 34/2022 adhesion test – testing of selected adhesives on fired clay references [1] w. brameshuber, m. graubohm. prefabricated masonry panel system with two-component polyurethane adhesive / vorgefertigte mauertafeln mit zweikomponenten-polyurethanklebstoff. cement and concrete composites 19:3–26, 2015. https://doi.org/10.1002/dama.201500644. [2] k. richterová. precast masonry wall panels. master thesis 2021. [3] heluz. heluz sidi. https://www.heluz.cz/sidi/. [4] heluz. heluz sb malta pro tenkou spáru. https://www.heluz.cz/files/sb_sbc.pdf. [5] heluz. heluz pěna (tenkovrstvé lepidlo). https://www.heluz.cz/files/heluz-trend_m5_p_na_ technicky-list_cz.pdf. [6] d. braven. multi kleber. https://www.denbraven.cz/produkt/multi-kleber/. [7] d. braven. nízkoexpanzní pěna. https: //www.denbraven.cz/produkt/nizkoexpanzni-pena/. [8] d. braven. mamut glue high tack. den braven. https://www.denbraven.cz/produkt/ mamut-glue-high-tack/. [9] illbruck. ot120 univerzální chemická kotva. https://www.illbruck.com/cs_cz/produkt/ ot120-univerzalni-chemicka-kotva/. [10] d. braven. lepidlo na dřevo woodfix d3. den braven. https://www.denbraven.cz/produkt/ lepidlo-na-drevo-woodfix-d3/. 93 https://doi.org/10.1002/dama.201500644 https://www.heluz.cz/sidi/ https://www.heluz.cz/files/sb_sbc.pdf https://www.heluz.cz/files/heluz-trend_m5_p_na_technicky-list_cz.pdf https://www.heluz.cz/files/heluz-trend_m5_p_na_technicky-list_cz.pdf https://www.denbraven.cz/produkt/multi-kleber/ https://www.denbraven.cz/produkt/nizkoexpanzni-pena/ https://www.denbraven.cz/produkt/nizkoexpanzni-pena/ https://www.denbraven.cz/produkt/mamut-glue-high-tack/ https://www.denbraven.cz/produkt/mamut-glue-high-tack/ https://www.illbruck.com/cs_cz/produkt/ot120-univerzalni-chemicka-kotva/ https://www.illbruck.com/cs_cz/produkt/ot120-univerzalni-chemicka-kotva/ https://www.denbraven.cz/produkt/lepidlo-na-drevo-woodfix-d3/ https://www.denbraven.cz/produkt/lepidlo-na-drevo-woodfix-d3/ acta polytechnica ctu proceedings 34:85–93, 2022 1 introduction 2 materials and samples 3 adhesion measurement 4 results and discussion 5 conclusion list of symbols acknowledgements references acta polytechnica doi:10.14311/app.2020.27.0090 acta polytechnica 27(0):90–95, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app application of steel fibres in alkali-activated mortars götz hüskena, ∗, lars-christian wagnera, b, gregor j. g. glutha, stephan pirskawetza, hans-carsten kühnea a federal institute for materials research and testing (bam), department of safety of structures (dpt.7), unter den eichen 87, 12205 berlin, germany b technische universität berlin, group of building materials and construction chemistry, gustav-meyer-allee 25, 13355 berlin, germany ∗ corresponding author: goetz.huesken@bam.de abstract. alkali-activated materials are ideal for the repair of concrete structures in harsh environmental conditions due to their high durability in chemically aggressive environments. however, slag-based mortars, in particular, are prone to shrinkage and associated cracks. in this respect, the application of steel fibres is one solution to reduce the formation of shrinkage induced cracks and to improve post cracking behaviour of these mortars. this study investigated the influence of two different types of steel fibres on the tensile properties of two alkali-activated mortars. direct tensile tests and single fibre pull-outs were performed to analyse the determining failure modes both on macro and micro scale. mechanical testing was accompanied by non-destructive testing methods such as digital image correlation and acoustic emission for a detailed analysis of the fracture process. keywords: alkali-activated materials, fibre pull-out, steel fibres, tensile strength. 1. introduction alkali-activated materials (aams) have been an extensively investigated topic in material sciences as they are a suitable alternative to systems based on ordinary portland cement (opc), in certain applications. much of the research covers design aspects using different precursor materials (such as fly ash or slag) and suitable alkaline activators (for example sodium silicates or sodium hydroxide) and how these variables influence mechanical and durability properties. these investigations show that aam, depending on their varying composition, can meet or even surpass properties of opc concretes [1–3]. nevertheless, aam often tend to be prone to higher shrinkage and more brittle behaviour than opc concretes [4]. fibres, such as steel, glass, synthetic, and natural fibres, within the concrete matrix have often been used to increase ductility and to reduce the formation of shrinkage induced cracks. adding fibres to the brittle matrix improves the post-cracking behaviour of the composite material significantly as the fibres connect the crack edges after cracks are formed. the application of fibres is usually limited to plastic fibres to reduce effects resulting from the formation of shrinkage cracks, but steel fibres are also of special interest for the application in aam [5]. besides reducing the formation of shrinkage induced cracks, the bond behaviour between steel fibres and the surrounding aam matrix is in some cases stronger than in opc concretes [6]. however, this effect is not yet fully understood. the application of steel fibres up to 2 % to 3 % by volume significantly increases both flexural behaviour and fracture energy of the aam composite, which is mainly caused by the high tensile strength and ductility of the fibres [7]. single fibre pull-out tests are a suitable test method to characterize the bond strength between steel fibres and the surrounding matrix [8]. these tests are common for opc steel fibre-reinforced concretes but are seldom used for characterizing the fracture behaviour of aam. steel fibre-reinforced aams prove to be alternatives to steel fibre-reinforced opc in fields like repair tasks, underground engineering or pre-cast elements [9, 10]. it is thus important to investigate the failure mechanisms of steel fibre-reinforced aam regarding: (i)the influence of the fibre properties such as length, diameter and form on the mechanical properties of aams. (ii)the influence of the (chemical) composition of the aams on the bond between the matrix and the fibres’ surfaces. (iii)how different compositions may lead to different failure modes and post-cracking behaviour. 2. materials and experimental setup two types of aam based on ground-granulated blastfurnace slag (ggbs) and fly ash have been investigated in combination with two different steel fibres. the mortars are referred to as c7a (slag) and fa5 (fly ash), respectively. the composition of the investigated mortars is given in table 1. further details on 90 http://dx.doi.org/10.14311/app.2020.27.0090 http://ojs.cvut.cz/ojs/index.php/app vol. 27/2020 application of steel fibres in alkali-activated mortars the mortars can be found in [11, 12]. the properties of the applied steel fibres are summarized in table 2. material c7a (kg/m3) fa5 (kg/m3) slag/fly ash 574.3 579.7 activator 1 50.8 234.0 activator 2 89.0 71.4 water 176.1 13.3 sand 0.1-0.5 400.6 402.0 sand 0.5-1 328.1 239.4 sand 1-2 374.2 239.4 sand 2-4 220.3 452.1 activator 1 sodium silicate solution with a modulus of 2.1 (‘grade d’) activator 2 naoh solution. table 1. mortar composition. property fibre 1 fibre 2 length (mm) 13 20 diameter (mm) 0.2 0.3 l/d ratio 65 66.7 surface (mm2) 8.17 18.85 tensile strength (mpa) >2 600 >1 200 max. tensile force (n) 81.68 84.82 density (kg/m3) 7 850 7 850 fibres per kg 314 000 90 110 coating brass none table 2. properties of steel fibres used. mixes with a steel fibre content (fibre type 1) of 1 %, 2 % and 3 % by volume were tested for the mortar c7a and 2 % by volume for mortar fa5. the steel fibres of type 2 were added to mortar c7a in such a quantity that their number per volume is comparable to the mix with 1 % by volume (c7a_1%_f1). the direct tensile tests were performed according to the procedure described in [13]. the 56 days old specimens were tested in a displacement controlled uniaxial tension test with a loading rate of 0.0025 mm/s until a total displacement of 1.5 mm was reached. to ensure uniform crack opening, both ends of the specimens were glued in the sample holders fixed to the servo-hydraulic testing machine (see figure 1). two linear variable differential transformers (lvdts) were clamped laterally to the specimen to measure the deformation in the zone with constant cross section. the machine data (force and displacement) were recorded at a rate of 50 hz. the tensile tests were accompanied by digital image correlation (dic) and acoustic emission (ae) analysis. a high-contrast stochastic pattern was applied on the front side of the specimen for the dic measurements. the optical deformation measurement was performed using an aramis 5m system supplied by gom, gesellschaft für optische messtechnik mbh. ae-sensors clamp dic field lvdt figure 1. test setup. the dic system was either triggered by the testing machine and a signal was sent every 0.01 mm of piston displacement to the dic camera or a fast measurement with 20 frames per second was conducted. 4 acoustic emission sensors (2 vs150ms and 2 pac wd) were fixed to the specimen and acoustically coupled using viscoelastic adhesive pads bosticprestik®. the acoustic emission signals were processed and stored by an amsy6 acoustic emission system supplied by vallen systeme gmbh. the detection threshold was set to 37.1 dbae and the bandpass filter was set to 25 khz to 800 khz for all sensors. single fibre pull-out tests were conducted for all fibre-matrix combinations to characterize acoustic emission signals in terms of formation of micro cracks and fibre pull-out. the pull-out tests were performed displacement controlled with a loading rate of 0.0025 mm/s until a pull-out length of 0.5 mm was reached and was then increased to 0.0333 mm/s until total pull-out of the fibre. a detailed description of the sample preparation and further test details can be found in [14]. 3. results and discussion figure 2 shows the averaged stress-strain plots (3 specimens for each mix) for the tested steel fibre combinations. the loading response of the tested mortars can be divided into three characteristic segments. firstly, all samples show a linear-elastic behaviour until the maximum tensile stress (crack stress) is reached. this maximum tensile stress varies between 2.8 mpa to 91 g. hüsken, l.-c. wagner, g. j. g. gluth et al. acta polytechnica 0 1 2 3 4 5 6 7 0 1 2 3 4 5 6 7 t en si le s tr es s (m p a) strain (‰) c7a_1%_f1 c7a_2%_f1 c7a_3%_f1 c7a_3.5%_f2 fa5_2%_f1 figure 2. averaged stress-strain curves. 3.2 mpa for the mortar c7a. fa5 has a stronger matrix and shows with 6.3 mpa a maximum value that is about twice that observed for the tensile strength of c7a. further loading beyond peak results either in a major drop of the tensile stress that stabilizes at a constant low level for low fibre contents or the gradual decrease of the tensile stress for higher fibre contents. the performance of the tested mortars under uniaxial tensile loading can be described according to [15] as strain-softening behaviour (level 1). furthermore, the low deviation of the crack stresses of mortar c7a for all tested fibre contents demonstrates that incorporation of fibres does not increase the tensile strength of the material, but influences the post-critical behaviour beyond matrix cracking due to crack bridging [15]. here, not only the fibre content is an important factor, but also the bond behaviour between the fibre and the surrounding matrix is essential. this fact is expressed by the area under the stress-strain plots beyond peak for mortars with same fibre content (2 % by volume) but different matrix material (c7a and fa5). in this case, the material with higher tensile strength (fa5) seems to have also the better bond between fibre and matrix. further insights into the post-critical behaviour of this composite material can be derived from the results of ae analysis and dic measurements. figure 3 shows as example the results of the ae analysis of specimen fa5_2%_f1_3. in this diagram, the localized acoustic events and the applied tensile force are plotted versus time. furthermore, a picture of the dic measurement is added showing the strain distribution of the cracked specimen at the end of the tensile test. the localized acoustic events have been separated into two groups considering the weighted peak frequency (wpf) as suggested by [16]. the threshold of the wpf was set to 110 khz. this value is based on preliminary tests where a clustering of acoustic events at 70 khz and 150 khz were observed. 0 1 2 3 4 5 6 f or ce ( kn ) wpf > 110 khz wpf < 110 khz st ep 1 96 0 0 100 200 300 400 500 600 time (s) -8 -6 -4 -2 0 2 4 6 8 p os it io n ( cm ) figure 3. results of the ae analysis for specimen fa5_2%_f1_3. only few acoustic events have been detected in the uncracked state before maximum tensile stress. this low ae activity is related to the formation of micro cracks and increases rapidly in the upper part of the specimen shortly before crack formation causes a sudden drop in force. after the first crack is formed, the upper part of the specimen shows a high ae activity (wpf ≤110 khz) that indicates the activation of steel fibres due to the formation of micro cracks. starting at 85 seconds, the ae activity increases in the lower part of the specimen, indicating the formation of a second crack. the ae activity near the upper first crack decreases significantly after the second crack is formed. ae events (wpf ≤110 khz), indicating the formation of micro cracks, were mainly detected in the lower part of the specimen until the end, but their number decreases and higher frequency events (wpf >110 khz) dominate the fracture process in the lower part. these higher frequency events indicate fibre pull-out near the second crack as it occurs during crack opening. the observations made by ae analysis regarding crack formation and crack opening are in good agreement with the results of the dic measurements. figure 4 shows the strain distribution of specimen fa5_2%_f1_3 for relevant loading levels (steps). additionally, the corresponding tensile stress (mpa) for each step is given in parenthesis in figure 4. the corresponding steps of the dic measurement are marked in figure 3 by two dashed black lines. the first dashed line near the peak covers the time interval for the formation of the first and second crack, which is represented in figure 4 by the steps 1070, 1071, 1087 and 1088. the second dashed line shows the time interval at step 1960 when mainly crack opening occurs. step 1070 shows the specimen shortly before the peak is reached. all other pictures taken before are similar and do not show any relevant strains. 92 vol. 27/2020 application of steel fibres in alkali-activated mortars step 1088 (5.0) µm/m 30000 24000 21000 18000 15000 12000 9000 6000 3000 0 27000 step 1070 (6.5) step 1071 (6.1) step 1960 (4.5)step 1087 (5.6) figure 4. strain distribution of sample fa5_2%_f1_3. step 1071 was taken 0.05 s later and shows the formation of the first crack indicated by the strain concentration in the upper part of the specimen. the crack is formed from the left side of the specimen to the middle and results in a drop of the tensile stress of about 0.5 mpa. step 1087 was taken 0.8 s later and shows just a minor increase of the strains in the upper part. however, a second crack was formed in the lower part of the specimen, growing from the right side to the middle. step 1088 was taken, again, 0.05 s later and shows the rapid opening of the second crack, whereas the upper crack shows no significant changes in its strain distribution. the formation of the second crack results in a further drop of the tensile stress of about 0.6 mpa. step 1960 shows the strain distribution of the specimen at higher loading level when the tensile stress decreased to 4.5 mpa. the strain distribution in the upper part increases slightly, indicating an opening of the upper crack, but no lateral expansion to the right. the crack opening of the second crack is more dominant as it becomes evident from the larger increase in the strain distribution. moreover, the lower crack runs through the entire front surface of the specimen. single fibre pull-out tests were conducted to get further insights into the fracture behaviour in the post-critical phase. figure 5 shows typical loaddisplacement plots for both steel fibre types embedded in mortar c7a. in the beginning, the pull-out force increases linearly up to the maximum pull-out force. after reaching the peak, the force drops rapidly until a stable plateau is reached, which is dominated by the friction between the steel fibre and the surrounding matrix. in this phase, the pull-out force usually 0 1 2 3 4 5 6 7 8 9 10 0 5 10 15 20 25 30 p u ll -o u t fo rc e (n ) displacement (mm) maximum pull-out force linear-elastic part friction during pull-out pull-out blocking before pull-out c7a_f2_4 c7a_f1_3 figure 5. load-displacement plots of steel fibres type 1 and 2 embedded in mortar c7a. reaches a stable value or decreases slightly as with increasing fibre pull-out the force transmitting contact area between steel fibre and surrounding matrix decreases [17]. the pull-out force increases short before the fibre is pulled-out. similar findings are reported by [15] and have been explained by the plastic deformation of the fibre end due to the cutting process. the deformation of the fibre ends provides a mechanical anchorage, which also may help increasing the pull-out resistance. furthermore, it is possible according to [15] that some matrix particles adhere to the fibre surface. the abrasion of these particles and their accumulation 93 g. hüsken, l.-c. wagner, g. j. g. gluth et al. acta polytechnica towards the fibre pull-out end might cause a wedge effect leading to higher pressure on the fibre and higher frictional bond. both effects could be confirmed for both steel fibre types by own sem images as depicted in figure 6 for the plastic deformation of the fibre end and figure 7 for adhering matrix particles. 100 µm figure 6. plastic deformation of the fibre end due to the cutting process (steel fibre type 1). 100 µm figure 7. adhering matrix particles (c7a) at the fibre end (steel fibre type 1). it was already assumed from the averaged stressstrain plots that the bond behaviour between fibre and surrounding matrix of mortar fa5 is better than for mortar c7a. this assumption was confirmed by the results of the single fibre pull-out tests depicted in figure 8 for steel fibres type 1. here, not only a higher maximum pull-out force was obtained, but also the friction dominated pull-out phase differs from the results obtained for mortar c7a. in this case, the pull-out force decreases clearly with increasing fibre pull-out as the force transmitting contact area max. pull-out force friction during pull-out pull-out blocking during pull-out 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0 10 20 30 40 50 60 70 80 p u ll -o u t fo rc e (n ) displacement (mm) fa5_f1_4 fa5_f1_2 fibre rupture figure 8. load-displacement plots of steel fibres type 1 embedded in mortar fa5. between steel fibre and surrounding matrix decreases. a slight increase in the pull-out force short before complete fibre pull-out was also recorded for steel fibres embedded in mortar fa5. however, fibre rupture occurred for both types of steel fibres at different levels (see figure 9). therefore, it was not possible to obtain comparable load-displacement plots for steel fibres type 2 as they all ruptured before the fibre was pulled-out from the matrix. 100 µm figure 9. ruptured fibre (steel fibre type 1) with adhering matrix particles (fa5). 4. conclusions uniaxial tensile tests and single fibre pull-out tests have been conducted to characterize the determining failure modes of fibre-reinforced aam. two different aam mortars have been investigated in combination with two types of steel fibres. the uniaxial tensile tests revealed that the post-critical behaviour is affected by the matrix properties (e.g. fibre-matrix bond), fibre 94 vol. 27/2020 application of steel fibres in alkali-activated mortars content and type, whereas the cracking stress was only affected by the matrix strength. single fibre pull-out tests confirmed the findings that the bond behaviour between embedded fibre and surrounding matrix is the main factor to control the post-critical behaviour of fibre-reinforced aam. the uniaxial tensile tests have been performed along with ae analysis and dic measurements. both techniques delivered valuable information on the determining failure modes in the post-critical phase. the results of the ae analysis demonstrated that it is possible to distinguish between the formation of micro cracks with low frequency and fibre pull-out with high frequency. furthermore, increased ae activity near the formed cracks was measured which correlates well with the findings of the dic measurements. the combination of both non-destructive test methods proved to be a useful tool for analysing failure mechanism of this composite material and optimizing material properties. acknowledgements this study was undertaken as part of the bleib project within the federal institute for materials research and testing (bam) and funded by the federal ministry for economic affairs and energy (bmwi), germany. the authors gratefully thank ms berta mota gassó for her assistance with sem imaging. references [1] j. provis, j. van deventer (eds.). alkali activated materials. springer-verlag gmbh, 2013. [2] j. c. petermann, a. saeed, m. i. hammons. alkali-activated geopolymers: a literature review. tech. rep., air force research laboratory materials and manufacturing directorate , 2010. doi:10.21236/ada559113. [3] k. arbi, m. nedeljković, y. zuo, g. ye. a review on the durability of alkali-activated fly ash/slag systems: advances, issues, and perspectives. industrial & engineering chemistry research 55(19):5439–5453, 2016. doi:10.1021/acs.iecr.6b00559. [4] j. r. yost, a. radlińska, s. ernst, et al. structural behavior of alkali activated fly ash concrete. part 2: structural testing and experimental findings. materials and structures 46(3):449–462, 2012. doi:10.1617/s11527-012-9985-0. [5] s. bernal, r. m. de gutiérrez, e. rodríguez, et al. mechanical behaviour of steel fibre-reinforced alkali activated slag concrete. materiales de construcción 59(293):53–62, 2009. doi:10.3989/mc.2009.41807. [6] t. ng, t. htut, s. j. foster. mode i and ii fracture behaviour of steel fibre reinforced high strength geopolymer concrete: an experimental investigation. in proceedings of framcos-7. 2010. [7] s. s. kumar, k. c. pazhani, k. ravisankar. fracture behaviour of fibre reinforced geopolymer concrete. current science 113(01):116–122, 2017. doi:10.18520/cs/v113/i01/116-122. [8] k. wille, a. e. naaman. bond stress-slip behavior of steel fibers embedded in ultra high performance concrete. in proceedings of 18th european conference on fracture and damage of advanced fiber-reinforced cement-based materials, pp. 99–111. 2010. [9] f. pacheco-torgal, j. b. de aguiar, y. ding, et al. performance of alkali-activated mortars for the repair and strengthening of opc concrete. in handbook of alkali-activated cements, mortars and concretes, pp. 627–641. elsevier, 2015. doi:10.1533/9781782422884.4.627. [10] a. buchwald. what are geopolymers? current state of research and technology, the opportunities they offer, and their significance for the precast industry. bft international 72(7):42–49, 2006. [11] n. hu, g. hüsken, g. j. g. gluth, h.-c. kühne. the influence of activator composition on the strength, shrinkage and chloride migration resistance of alkali-activated slag mortars. key engineering materials 761:61–64, 2018. doi:10.4028/www.scientific.net/kem.761.61. [12] g. j. g. gluth, w. d. a. rickard, s. werner, s. pirskawetz. acoustic emission and microstructural changes in fly ash geopolymer concretes exposed to simulated fire. materials and structures 49(12):5243– 5254, 2016. doi:10.1617/s11527-016-0857-x. [13] s. m. pirskawetz, g. hüsken, i. curosu, v. mechtcherine. combination of digital image correlation and acoustic emission for characterizing failure modes in strain-hardening cement-based composites (shcc). in strain-hardening cement-based composites, pp. 300–307. springer netherlands, 2017. doi:10.1007/978-94-024-1194-2_35. [14] l.-c. wagner. untersuchung des einflusses von stahlfasern auf das bruchverhalten alkali-aktivierter mörtel. master’s thesis, technische universität berlin, berlin, germany, 2017. [15] k. wille, s. el-tawil, a. naaman. properties of strain hardening ultra high performance fiber reinforced concrete (uhp-frc) under direct tensile loading. cement and concrete composites 48:53–66, 2014. doi:10.1016/j.cemconcomp.2013.12.015. [16] m. sause, t. müller, a. horoschenkoff, s. horn. quantification of failure mechanisms in mode-i loading of fiber reinforced plastics utilizing acoustic emission analysis. composites science and technology 72(2):167– 174, 2012. doi:10.1016/j.compscitech.2011.10.013. [17] a. khabaz. analysis of sliding mechanism of straight steel fibers in concrete and determine the effect of friction. archives of civil and mechanical engineering 17(3):599–608, 2017. doi:10.1016/j.acme.2017.01.005. 95 http://dx.doi.org/10.21236/ada559113 http://dx.doi.org/10.1021/acs.iecr.6b00559 http://dx.doi.org/10.1617/s11527-012-9985-0 http://dx.doi.org/10.3989/mc.2009.41807 http://dx.doi.org/10.18520/cs/v113/i01/116-122 http://dx.doi.org/10.1533/9781782422884.4.627 http://dx.doi.org/10.4028/www.scientific.net/kem.761.61 http://dx.doi.org/10.1617/s11527-016-0857-x http://dx.doi.org/10.1007/978-94-024-1194-2_35 http://dx.doi.org/10.1016/j.cemconcomp.2013.12.015 http://dx.doi.org/10.1016/j.compscitech.2011.10.013 http://dx.doi.org/10.1016/j.acme.2017.01.005 acta polytechnica 27(0):90–95, 2020 1 introduction 2 materials and experimental setup 3 results and discussion 4 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0207 acta polytechnica ctu proceedings 33:207–211, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague equivalent thermal diffusivity of natural and recycled aggregate concrete at temperature up to 350 ◦c izabela hagera, ∗, katarzyna mróza, tomasz tracza, václav kočíb, lukáš fialab a chair of building materials engineering, faculty of civil engineering, cracow university of technology, 24 warszawska st. 31-155 cracow, poland b czech technical university in prague, department of materials engineering and chemistry, faculty of civil engineering, thakurova 7/2077, 16629 prague 6, czech republic ∗ corresponding author: izabela.hager@pk.edu.pl abstract. the paper presents the evolution of equivalent thermal diffusivity with temperature of concretes with natural coarse aggregates and recycled concrete aggregates. thermal diffusivity was evaluated based on the temperature measurements made on cylindrical samples using so-called "inverse technique". thermal diffusivity parameter was estimated for concretes manufactured with different replacement ratios of natural by recycled coarse aggregate (0, 50 and 100 %) produced by crushing the worn-out concrete pavement that was dismounted after exploitation period. the diffusivity of concrete was tested up to 350 ◦c. the progressive decrease of thermal diffusivity was observed, and the lowest d values were obtained for higher recycled concrete aggregates content. the results were presented against the background of other authors results and referred to the d values proposed in eurocode 2. keywords: fire, recycled concrete aggregates, temperature, thermal diffusion. 1. introduction by volume,construction and demolition waste (cdw) are one of the largest streams of waste in europe with 850 million tons that are produced per year [1]. the waste separation process enables to obtain good quality aggregates and their use in concrete is economically, environmentally and technically feasible with the respect of all technical standards. the progressive replacement of natural aggregates by recycled concrete aggregates contributes to the implementation of the strategy for sustainable development of the field of construction through the successful implementation of sustainable building materials with a reduced carbon footprint. the concrete industry scepticism based on concern about poor-quality recycled products needs to change with the research projects showing successful application of recycled concrete aggregates rca without compromising on quality and contribute to the decrease environmental impact. the detailed recommendations concerning assessing the properties of recycled aggregates are given in [2], however the imitation concerning their replacement of natural aggregate in the concrete mix is imposed [3]. in the present research the properties of concrete with recycled aggregates were investigated in the view of study their behaviour in fire. the fire behaviour of concrete with natural aggregate is well documented in the literature [4–6]. well known cases of fires in engineering facilities like channel (1996); mont blanc (1999); gotthard (2001) tunnels has drew attention to one of the disadvantages of concrete that is concrete spalling in a fire. the sudden loss of concrete cover and concrete element cross-section reduction may jeopardise the overall load-bearing capacity of the structural element [7]. it has been confirmed that the composition of concrete, as well as boundary conditions, affects the concrete propensity to spall in a fire. one of the main parameters governing concrete spalling is permeability. in denser and less permeable concrete, the spalling risk is higher [8]. concrete permeability is a derivative of the type of aggregate used, transport properties of the cement paste and the permeability of interfacial transition zone itz between the cement paste and the aggregate. the permeability of recycled concrete aggregate differs from natural aggregate, thus transport properties of those concretes may be higher than in traditional concretes [9], thus lowering the risk of spalling behaviour. in view of studying the rca concrete behaviour in the fire the first step that was undertaking was the observation of thermal properties evolution presented in this paper. the aim of this project is to examine the impact of concrete composition, in particular the recycled concrete aggregate amount rca on the thermal properties of concrete. as the heat transfer in concrete plays a crucial role in fire resistance evaluation, the 207 https://doi.org/10.14311/app.2022.33.0207 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en i. hager, k. mróz, t. tracz et al. acta polytechnica ctu proceedings figure 1. thermocouples arrangement in the concrete cylinder, and the furnace heating rate. thermal properties that change with the temperature are essential and need to be provided to the calculation model. thermal diffusivity (d) is a specific property of the material and allows you to determine how quickly a given material reacts to temperature changes. this is a property that measures the rate of heat transfer from the exposed surface of the material to the inner layers. the higher the value of thermal diffusivity, the faster the temperature increase at a certain depth. to predict the behaviour of concrete during a fire and carry out simulations determining the temperature development in a heated element, one need to know the value of thermal diffusivity. thermal diffusivity can be evaluated directly from its definition or its evaluation can be based on the temperature measurements made on cylindrical samples, called an "inverse technique". 2. experimental method the thermal diffusivity (d) of concrete depends on its thermal conductivity (λ), density (ρ) and specific heat of the material (cp). thermal diffusivity can be determined (eq. 1) by directly determining these physical properties by experimental methods or by using an inverse method leading to the determination of d. the unit of thermal diffusivity is [j/kgk], but [mm2/s] or [m2/s] are also used in the literature. d = λ ρ cp ! m2 s " (1) the method determining the equivalent thermal diffusivity of concrete was recently described by felicetti [10, 11] showing good compliance with experimental results from standard fire tests. the main advantage of the "inverse technique" is the determination of thermal diffusivity without the need for cumbersome measurements of thermal conductivity and specific heat values as well as density change evaluation, knowing that all those values vary with temperature increase. the inverse method is quite sensitive to moisture presence in concrete and other physicochemical processes that occur when the material is exposed to high temperature. it consists in heating the cylindrical sample (diameter 100 mm; height 200 mm) in the electric furnace and measuring temperature increase in the centre of the sample (tc centre) and on 2 mm from cylinder lateral surface (te exterior) using k-type thermocouples, figure 1. the temperature differences are measured on the central crosssection where the field of temperature is considered dependant only to distance from the cylinder centre. the temperature increases with the constant heating rate 1 ◦c/min up to 350 ◦c. the test result is the temperature difference te − tc in function of time. during the test. in the determination of thermal diffusivity by the inverse method, the initial assumptions is adopted that the cylinder is axis-symmetric and infinitely long, so the analyse is related to a differential equation for transient heat conduction. to determine the diffusivity that characterises the transient process of heat exchange inside a constant mass we use fourier’s law (eq. 2): ρ cp ∂t ∂t + div (q) = 0 (2) assuming a cylindrical coordinate system (θ, r, z) (eq. 3): ∂t ∂t = d ! d2t dr2 + 1 r ∂t ∂t " (3) and using laplace transformation, we get a formula for the temperature difference between the edge and the centre of the sample (eq. 4): 208 vol. 33/2022 natural and recycled aggregate concrete thermal diffusivity component units oc rac 50% rac 100% cem i 42.5 r kg/m3 414.2 414.2 414.2 water kg/m3 186.4 186.4 186.4 w/c − 0.45 0.45 0.45 river sand 0/4 mm dwudniaki kg/m3 615.1 280.6 569.2 coarse agg. 2/8 mm dwudniaki kg/m3 579.9 280.6 − coarse agg. 8/16 mm dwudniaki kg/m3 562.3 264.1 − recycled concrete agg. 4/16 mm kg/m3 − 825.3 1057.2 plasticizer basf bv 18 % mc 0.9 0.9 0.9 basf glenium sky 591 % mc 1.4 1.4 1.4 cement paste content dm3/m3 320 320 320 mortar content dm3/m3 542 511 540 density kg/m3 2248.5 2236.8 2221.3 water content % 5.2 5.8 6.7 water absorption % 6.9 7.3 7.8 28 days compressive strength mpa 46.1 56.8 54.5 28 days tensile splitting strength mpa 3.3 3.6 3.4 table 1. mix design and properties. t (r, t) = r ! t − a2 − r2 4d " (4) from here we get the formula for d equivalent thermal diffusivity (eq. 5): d = ra2 4∆t (5) where: • r heating rate [◦c/min], • a sample radius [m], • ∆t temperature gradient [◦c/cm], • d -thermal diffusivity [mm2/s]. 3. materials the measurements were carried out for concrete made of cem i 42.5 cement, riverbed coarse aggregate dwudniaki from dunajec river southern poland (grain size 2/8 mm and 8/16 mm) and crushed concrete from the worn-out concrete pavement with grain size 4/16 mm. before being crushed the mechanical properties of concrete from the pavement was determined on drills. the compressive strength of 66.8 mpa, modulus of elasticity 41.2 gpa and apparent density of 2290 kg/m3 characterised the recycled concrete. for all 3 concretes the same water-cement ratio (w/c) and cement paste volume was maintained. additionally, a plasticiser was used to obtain good workability and water content reduction. the mix composition of the concretes is shown in table 1. the concretes were labelled: the reference ordinary concrete (oc) made of natural riverbed aggregates; rac 50% is concrete made of recycled pavement aggregate with 50% of natural coarse aggregate replacement with recycled one; rac 100% recycled aggregate concrete made of recycled concrete coarse aggregates. in all concretes oc, rac 50% and rac 100% riverbed quartz sand was used. the concretes specimens were cast and stored in plastic cubic or cylindrical moulds for the first 24 hours. after preliminary curing, water evaporation was prevented by covering them with plastic lids for 7 days. during the next 28 days, they were stored in natural air-drying conditions at the temperature of 20 ± 5 ◦c and relative humidity of 50 ± 5 %. the compressive strength and splitting tensile strength of concretes after 28 days were determined on 10 cm3 cubes. all basic physical (density, water content, water absorption) and mechanical (compressive and splitting tensile strength) performances are summarised in table 1. 4. results and observations the temperature evolution in time measured with type k thermocouples was presented in figure 2. the solid line is showing the heating programme. the presented results were used to determine effective thermal diffusivity as a function of temperature, obtained from the eq. 5 and expressed in [mm2/s]. the results are shown in figure 3. due to thermal gradient value close to zero at the beginning of the test, when the furnace starts to heat, the effective thermal diffusivity values are very high. this phenomenon was previously observed by [10], and the correction factor was proposed to reduce this unwanted effect so that the suggestion to compute the instantaneous heating rate was proposed [11]. the equivalent thermal diffusivity decreases with increasing temperature in the range up to 350 ◦c. in heated 209 i. hager, k. mróz, t. tracz et al. acta polytechnica ctu proceedings figure 2. time temperature curves for oc, rac 50% and rac 100%. figure 3. equivalent thermal diffusivity for oc, rac 50% and rac 100%compared with the results of felicetti [11] and ec 2 proposed values. with 1 ◦c/min cylinders the values of equivalent thermal diffusivity reaches at 100 ◦c 0.76 mm2/s, 0.61 mm2/s and 0.45 mm2/s respectively for oc, rac 50% and rac 100%. the highest value obtained for oc results from the density of this material. with an increasing proportion of recycled concrete aggregates, the d values are reduced. the similar tendencies are observed at 200 ◦c and 300 ◦c. from the graph, we can see the water migration effect occurring in the lower temperature in case of concretes with recycled concrete aggregates. it confirms that those materials enable faster water migration due to their higher permeability. the observed water migration perturbance on figure 3 is linked with specific heat significant peak at a temperature close to 200 ◦c. the comparison with the d range proposed by eurocode 2 (ec2) [12] for both samples with recycled aggregates rac 50% and rac 100% shows thatthe thermal diffusivity values are mostly below both the upper and lower limits given in ec2. the results remain similar to the experimental results of felicetti [10, 11]. 5. conclusions determination of thermal diffusivity is a vital parameter for the correct modelling of heat flow through heated concrete. thanks to the inverse method, we determine diffusivity for much higher temperatures than other stationary and non-stationary methods. non-stationary transient temperature methods, despite their broad application, require the use of specialised and often costly measuring equipment complex in mathematical and numerical terms. in the inverse method, we can use widely available and easyto-use laboratory equipment to perform the tests. although the technique does not allow for specific parameters to established: ρ, cp and λ but their ratio 210 vol. 33/2022 natural and recycled aggregate concrete thermal diffusivity expressed in diffusivity definition d = λ/ρ cpis assessed. acknowledgements this research was supported and funded in part from the emmat project e-mobility and sustainable materials and technologies ppi/apm/2018/1/00027 financed by the polish national agency for academic exchange (nawa). references [1] global recycling. recycled construction waste in europe. https://global-recycling.info/archives/2964. [2] rilem tc 121-drg. specifications for concrete with recycled aggregates. materials and structures 27(173):557-59, 1994. [3] en 12620:2004 + a1:2010. aggregates for concrete, 2004. [4] j.-c. mindeguia, p. pimienta, h. carré, et al. on the influence of aggregate nature on concrete behaviour at high temperature. european journal of environmental and civil engineering 16(2):236-53, 2012. https://doi.org/10.1080/19648189.2012.667682. [5] z. xing, a.-l. beaucour, r. hebert, et al. influence of the nature of aggregates on the behaviour of concrete subjected to elevated temperature. cement and concrete research 41(4):392-402, 2011. https: //doi.org/10.1016/j.cemconres.2011.01.005. [6] i. hager, t. tracz, j. śliwiński, et al. the influence of aggregate type on the physical and mechanical properties of high-performance concrete subjected to high temperature. fire and materials 40(5):668-82, 2016. https://doi.org/10.1002/fam.2318. [7] k. d. hertz. limits of spalling of fire-exposed concrete. fire safety journal 38(2):103-116, 2003. [8] i. hager, k. mróz. role of polypropylene fibres in concrete spalling risk mitigation in fire and test methods of fibres effectiveness evaluation. materials 12(23), 2019. https://doi.org/10.3390/ma12233869. [9] m. choinska, a. khelidj, g. chatzigeorgiou, et al. effects and interactions of temperature and stress-level related damage on permeability of concrete. cement and concrete research 37(1):79-88, 2007. https: //doi.org/10.1016/j.cemconres.2006.09.015. [10] r. felicetti. thermal diffusivity and residual strength of high-performance light-weight concrete exposed to high temperature. proc. 6th int. symposium on the utilisation of high-strength/high performance concrete 2, leipzig (germany) 935-48, 2002. [11] r. felicetti. assessment of the equivalent thermal diffusivity for fire analysis of concrete structures, proc. fib task group 4.3 workshop fire design of concrete structures, coimbra 149-58, 2007. [12] en 1992-1-2. eurocode 2: design of concrete structures part 1-2: general rules structural fire design 2004, 1992. 211 https://global-recycling.info/archives/2964 https://doi.org/10.1080/19648189.2012.667682 https://doi.org/10.1016/j.cemconres.2011.01.005 https://doi.org/10.1002/fam.2318 https://doi.org/10.3390/ma12233869 https://doi.org/10.1016/j.cemconres.2006.09.015 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0026 acta polytechnica ctu proceedings 34:26–31, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague benchmark vercors 2022: mechanical response of the prestressed concrete containment wall to ambient conditions štěpán krátký∗, petr havlásek czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: stepan.kratky@fsv.cvut.cz abstract. the vercors benchmark was developed to provide a solid experimental basis for numerical modeling of concrete containment buildings (ccbs) to extend their lifespan. the goal of the first phase of the third benchmark is to present a blind prediction of ccb’s behavior across the whole lifespan and under periodic pressure tests. this paper summarizes the calibration of the material model based on the microprestress-solidification (mps) theory according to the provided laboratory experimental data. a special emphasis is given to the influence of the ambient conditions on drying creep and transient thermal creep of concrete and the calibration of the corresponding material parameters. these phenomena are studied on a representative section of the containment wall using weakly coupled thermo-hygro-mechanical analyses. keywords: concrete, creep, shrinkage, drying, cyclic temperature, modeling. 1. introduction in the present days, the main portion of electricity in france (up to 75%) comes from nuclear power plants. nowadays, a considerable number of atomic reactors (a little less than 20%) approach the end of their design lifespan. if this matter is not properly addressed in the near future, it might result in a vast energy shortage in france, not mentioning the increased risk of accidents. the main factor of the limited service life is the concrete containment building (ccb). when upgrading to more powerful 1300 mw reactors, electricité de france (edf) decided to enhance the safety of nuclear power plants and thereby modify the old ccb design. the new design (as shown in fig. 1) consists of two one-meter-thick prestressed concrete walls between which a constant vacuum slightly below the atmospheric pressure is maintained. such a measure ensures that no contamination escapes from the interior. to shut down old reactors and to build new ones proves to be enormously expensive, inconsiderate to the environment, and politically inadmissible. the other possibility is to extend the service life of existing containments by 20 years (from 40 to 60 years). that brings edf to a complicated task to accurately predict more than half a century of ccb’s lifespan. with the associated governing phenomena (creep and shrinkage of concrete and steel relaxation), such prediction becomes even more challenging. vercors mock-up is a diminished model of a ccb in a scale 1:3. according to the diffusion theory, the drying rate is proportional to the square of the representative depth. therefore, at the present scale, the drying process becomes accelerated 9× which means that the extended lifetime of 60 years can be analyzed and assessed in no more than 7 years. the vercors mock-up (fig. 1) was completed in 2014 and is 28 meters tall, 21 meters in diameter and with hundreds of sensors and thousands of meters of optical cables became the most studied concrete structure in the world. the wall of the inner containment is prestressed with steel tendons in both vertical and horizontal directions, which leads to the compression of 14 mpa and 8 mpa, respectively. every 12 months, the pressure test is conducted to verify the air tightness of the containment. in the ccb, the imposed 12-hour-long overpressure can reach up to 6 bar (0.6 mpa) needs to be considered. in addition to structural loading, the walls are exposed to varying ambient conditions. this paper is organized as follows. the first section summarizes the standard laboratory experiments which supplement the mock-up with suitable data for calibration of the constitutive models which are briefly outlined afterwards. next, with the identified material parameters, the representative periodic section of the inner containment wall is analyzed. this reduced model was created to cut the computational time of the full model and at the same time to easily verify the functionality of the most important features. a special attention is given to the sensitivity of the model to the variable ambient conditions and the associated material parameters. the response is evaluated in terms of the stresses and strains in the hoop (tangential) and vertical directions, and the evolution of the prestress. this article incorporates and extends the findings obtained within the final thesis of the first author [2]. 26 https://doi.org/10.14311/app.2022.34.0026 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 34/2022 benchmark vercors 2022: response to ambient conditions figure 1. section of the ccb vercors mock-up [1], dimensions in [mm]. 2. materials and methods 2.1. laboratory experiments the composition of the concrete used in the mock-up and laboratory experiments is summarized in table 1 and is identical to the concrete of nogent sur seine nuclear power plant. the high value of water-toingredients qtty [kg/m3] cement cem i 52.5n 320 water 167.9 aggregates (fine 830, coarse 1005) 1835 superplasticizer 2.6 table 1. concrete composition [3]. cement ratio, w/c = 0.525, predestines this concrete to a larger magnitude of drying shrinkage and creep. apart from the conventional short-term measurements at the age of 28 days (fcm = 48.7 mpa, e = 34.3 gpa, ft = 4.4 mpa), the laboratory data set comprised the following experiments: • basic creep and autogenous shrinkage • total creep and drying shrinkage at henv = 0.5 • moisture loss • porosity and aging sorption isotherm every experiment was conducted under both room (20◦c) and elevated temperature (40◦c). 2.2. material models 2.2.1. heat and moisture transport in the present study, the heat and moisture transport are for simplicity treated as independent processes without cross-coupling. heat conduction is idealized by using a linear transport model characterized by its heat conductivity and capacity. concrete drying is described by a widely accepted model proposed by bažant and najjar [4]. under the assumption of linear desorption isotherm, the governing equation for the diffusion of water vapor reads ∂h ∂t = ∇· (c(h)∇h) (1) where ∇h is the gradient of relative humidity and c(h) is the humidity-dependent diffusivity. for cementitious materials, this dependence is highly nonlinear and can be approximated as c(h) = c1  α0 + 1 −α0 1 + ( 1−h 1−hc )n   (2) where c1 is the maximum diffusivity at h = 1, α0 determines the ratio between minimum diffusivity at h = 0 and c1, and parameters hc and n describe the relative humidity threshold and the steepness of the transition. the ambient relative humidity is prescribed using a mixed boundary condition which relates the humidity flux jh with the humidity difference at the boundary via surface factor f. 2.2.2. structural analysis the mechanical behavior of concrete is described by a modified constitutive model based on the microprestress-solidification (mps) theory [5]. under sealed conditions and constant room temperature, the behavior is defined by the basic creep compliance function of the b3 model [6] which is entirely captured by 4 parameters q1 −q4. yet, in concrete, changes both in relative humidity and temperature give rise not only to volume changes (shrinkage/swelling or thermal dilation), but also to further creep (pickett effect or transitional thermal creep). this additional creep is primarily controlled by the parameter k3 which is different from the original mps model and which has been introduced [7] to 27 štěpán krátký, petr havlásek acta polytechnica ctu proceedings minimize the size effect on drying creep. shrinkage strain and relative humidity are linearly linked via their rates, ε̇sh = kshḣ (3) where ksh is a material parameter usually treated as a humidityand age-independent constant. in the analysis, the stresses in steel reinforcement are far below its yield stress, therefore it is described by an isotropic linear elastic material with young’s modulus set to 200 gpa. the behavior of prestressing tendons is characterized by a generalized model from eurocode 2 for class 2 steel with reduced relaxation characteristic strength 1620 mpa and e = 190 gpa. the influence of elevated or variable temperature on the rate of relaxation and prestress losses due to friction or slip at the anchors are not considered. 2.3. computational models and calibration strategy the experimental data sets which served for the calibration of the material models for concrete (rheological and diffusion properties) were provided by edf at the beginning of the benchmark. at room temperature the experiments were done on cylinders 160×1000 mm while at the elevated temperature the dimensions were 100×200 mm. the experimental data had to be compensated for the spurious shrinkage strains caused by unintended moisture leaking. the basic creep parameters q1 −q4 of the b3 model were determined first. the deformation of the creep experiment which started at the age of 90 days was compensated for the autogenous shrinkage measured on the companion specimen. (since the measurements began at the age of 90 days, the recorded value and was very small, ≈ 50 × 10−6.) owing to the homogeneous stress state, the calibration was done using an analytical expression for the compliance function of the b3 model. the identified values which were slightly different from the prediction given by the b3 model are summarized in table 3. in the remaining experiments, the stress state in the specimen is no longer uniform, which explains why a more advanced computational model was necessary to be developed. the resulting axisymmetric models with suitable boundary conditions reflect the behavior of a thin representative section from the mid-height. in this region, the stress distribution can be considered as a function of the distance from the axis of the specimen and entirely free from the boundary effects. the numerical problem was solved using a staggered scheme. in every time step, the heat and moisture transport subproblems were solved first and subsequently were followed by the structural analysis which utilized the computed fields of temperature and relative humidity. calibration procedure continued with the parameter controlling the magnitude of drying shrinkage and the parameters of the bažant-najjar diffusion model. the suitable values were determined by hand fitting and the response of the models was checked against the evolution of drying shrinkage and moisture loss at room temperature. with these parameters set, the parameter k3 was identified from the drying creep experiment, and finally, the effect of elevated temperature was tuned by adjusting parameter kt c of the modified mps theory. the summary of all material parameters is listed in table 3. visual outputs of the calibration can be found in the final thesis of the first author [2]. parameter value q1 9.0 × 10−6 mpa−1 q2 70.0 × 10−6 mpa−1 q3 25.0 × 10−6 mpa−1 q4 6.0 × 10−6 mpa−1 ksh 1.0 × 10−3 k3 10 kt m 6.5 table 2. identified parameters of material model mps. parameter value c1 28.2 mm2day−1 α0 0.055 hc 0.7 n 10 f 1.08 mm day−1 table 3. identified parameters of material model bažant–najjar for moisture diffusion. figure 2. computational model of the representative section for the structural sub-problem. despite the impressive extent of the vercors experimental program, the laboratory data do not provide sufficient evidence for proper calibration of the material model under general ambient conditions. the obstacle is that the investigated inner containment of the mock-up is subject to cyclic temperature and humidity which produce an additional increase in the creep rate of concrete. in the literature, this subject is covered rather insufficiently, although the 28 vol. 34/2022 benchmark vercors 2022: response to ambient conditions scarce existing data indicate that the creep rate is most significant during the first cycle (both thermal and humidity) and with subsequent cycling gradually decreases. another computational model was created to investigate the influence of the cyclic ambient conditions on the behavior of the containment wall and to estimate a suitable value of the parameter kt c which is responsible for the damping of the transitional thermal creep under temperature cycles this model is a representative section (figure 2) of the inner containment wall. as shown in the figure, in addition to both vertical and horizontal prestressing cables, the model comprises conventional reinforcement. the dimensions of the computational model are determined by the spacing of the cables. in the tangential (horizontal) direction of the wall, the model corresponds to an angle of 2π/160 as there are 160 equally spaced vertical cables. the lateral sides are not parallel. the horizontal cables have two different radii, and for this reason the height of the model is double of the vertical spacing, 264 mm. in the transport subproblems the spatial discretization is refined towards the inner and outer surface. due to the zero flux in the vertical direction, only one element per height of the model is used. in the hoop direction, the discretization is uniform and is identical to the structural subproblem. the boundary conditions of the structural analysis restrain the displacement normal to the lateral sides of the model. in the vertical direction, the periodic conditions are defined which permits not only overall elongation or shortening, but also warping of the horizontal surfaces. with these boundary conditions, the model should realistically capture the behavior at mid-height of the containment inner wall where the influence of the stiffener at the top and the massive foundation at the bottom is negligible. the self-weight is not considered since the effect would be very subtle in contrast to vertical prestressing. the reinforcement is discretized by linear truss elements which are connected to the linear hexahedral fe mesh of concrete (shown in figure 2) by means of hanging nodes. the slip between steel and concrete can be neglected due to the rotational symmetry of the structure and the resulting uniformity of the loading. 3. results and discussion 3.1. overall behavior the figures presented in this section document the evolution of stresses or strains computed with different combinations of ambient conditions and/or material behavior of concrete and prestressing steel. the notation in the legends is consistent and is summarized in the following list. • jb + relax: basic creep of concrete (henv = 0.98), relaxation of prestressing steel, t = 20◦c throughout the simulation • jb + no relax: ditto + no steel relaxation • henv1: creep and shrinkage at constant ambient temperature t = 20◦c and humidity from table 4, steel relaxation • henvi + ti: creep and shrinkage at ambient conditions specified in tables 4 and 5, steel relaxation unaffected by temperature • + test: additional internal overpressure notation description henv1 constant after 250 days henv2 cyclic changes, piecewise constant henv3 cyclic changes, piecewise linear henv1 = 0.26 on the inner side and 0.58 on the outer side with cyclic periods of henv = 0.98 on both faces. table 4. definition of ambient relative humidity. notation description t1 constant after 90 days t2 cyclic changes, piecewise linear t = 28◦c on the inner side and 25◦c on the outer side with cyclic periods of t = 10◦c on both faces. table 5. definition of ambient temperature. in the following figures, time t = 0 corresponds to the concrete age of 90 days. until t = 45 days when the prestressing is applied, the structural model of ccb wall is subject only to ambient conditions as specified by the keyword in the legend. figure 3 shows the evolution of strain in the hoop and vertical directions computed with different setup of the computational model. the results suggest that drying plays a fundamental role in the behavior of the containment wall. in the hoop direction, the strain after 2500 days of drying doubles the response of the models with basic creep only (red and purple curves), in the vertical direction the increase is even higher. temperature cycling (blue and green curves) causes a dramatic increase in strain whose rate is almost linear during the first 4 years. the computed evolution of the hoop stress in the middle and close to the inner surface is shown in fig. 4. prior to prestressing, the hoop stress close to the inner surface (fig. 4 right) induced by the drying shrinkage reaches almost 5 mpa. such a high tensile stress would have caused cracking which is not captured by the present material model. however, the crack depth would have been insignificant in contrast to the wall thickness and in addition to this, soon afterwards the containment wall becomes prestressed in both vertical and hoop directions which restores compression in the entire cross-section. the hoop stress in the middle of the wall is induced primarily by prestressing and only partially by the internally restrained drying shrinkage and nonuniform thermal strains. therefore, the decrease in stress magnitude in the middle of the wall can be attributed to the relaxation of the prestress caused by both creep 29 štěpán krátký, petr havlásek acta polytechnica ctu proceedings -300 -250 -200 -150 -100 -50 0 50 0 500 1000 1500 2000 2500 ta n g e n ti a l s tr a in , ε q [ 1 0 -6 ] time, t [day] jb + relax henv1 henv1 + t1 jb + no relax henv3 + t2 henv2 + t2 -150 -120 -90 -60 -30 0 30 0 500 1000 1500 2000 2500 v e rt ic a l s tr a in , ε z [ 1 0 -6 ] time, t [day] jb + relax henv1 henv1 + t1 jb + no relax henv3 + t2 henv2 + t2 figure 3. evolution of hoop (left) and vertical (right) strain in the center of the wall computed for different histories of ambient conditions, with basic creep only and with no steel relaxation. -15 -10 -5 0 0 500 1000 1500 2000 2500 s tr e s s, σ θ [ m p a ] time, t [day] jb + relax henv1 henv1 + t1 jb + no relax henv3 + t2 henv2 + t2 -15 -10 -5 0 5 10 0 500 1000 1500 2000 2500 s tr e s s, σ θ [ m p a ] time, t [day] jb + relax henv1 henv1 + t1 jb + no relax henv3 + t2 henv2 + t2 figure 4. evolution of hoop stress σθ in the middle (left) and close to the inner surface (right) of the ccb concrete wall. and shrinkage of concrete. steel relaxation leads to an initial yet considerable drop in prestressing which occurs during the first day. afterwards, this effect becomes negligible (see comparison in fig. 5). it is apparent that the highest decrease of prestress is associated with cyclic temperature and humidity loading (blue and green lines). 800 900 1000 1100 1200 1300 1400 1500 1 10 100 1000 10000 s tr e s s in t h e t e n d o n s , σ p [ m p a ] time, t-45 [day] jb + relax henv1 henv1 + t1 jb + no relax henv3 + t2 henv2 + t2 figure 5. decrease in prestress in horizontal tendons due to steel relaxation, concrete creep and shrinkage. 3.2. response to ambient conditions and pressure tests -300 -250 -200 -150 -100 -50 0 50 0 500 1000 1500 2000 2500 ta n g e n ti a l s tr a in , ε q [ 1 0 -6 ] time, t [day] henv1 + t1 henv2 t2 t2, ktc = ktm/20 figure 6. calibration of material parameter ktc responsible for the cyclic transient thermal creep. the analysis was performed with different histories of the ambient conditions which differed by the degree of simplification. the aim was to estimate the influence of this simplification which can significantly reduce the computational time. the results unanimously show that the temperature defined by a cyclic history t2 gives rise to an enormous increase in compliance which is only slightly affected 30 vol. 34/2022 benchmark vercors 2022: response to ambient conditions by the chosen history of relative humidity. having identified the source of this unrealistic response, a parameter kt c was introduced and based on previous experience set to kt m/20. with this approach, the response of the computational model can be damped in further temperature cycles provided that the temperature does not exceed its maximum previously attained value. the responses obtained with the original and modified formulation are shown in black and green color in fig. 6. -300 -250 -200 -150 -100 -50 0 50 0 500 1000 1500 2000 2500 ta n g e n ti a l s tr a in , ε q [ 1 0 -6 ] time, t [day] jb + relax henv3 + t2 henv2 + t2 jb + test henv1 + t1 henv1 + t1 + test figure 7. evolution of hoop strain in the center of the wall pressure tests. -15 -10 -5 0 0 500 1000 1500 2000 2500 s tr e ss , σ θ [m pa ] time, t [day] jb + relax henv3 + t2 henv2 + t2 jb + test henv1 + t1 henv1 + t1 + test figure 8. evolution of hoop stress in the center of the wall pressure tests. response to regular pressure tests is presented in figure 7 and 8. with the simplified model, the overpressure affects only the hoop stress which rises by ≈ 8.5 mpa, see figure 8. in contrast to the sustained loading by prestressing and drying, the pressure tests provide an interesting insight into the evolution of the incremental stiffness which increases with concrete aging and decreases with temperature cycles. 4. conclusions this paper briefly outlined a procedure used for the calibration of constitutive models for the rheological properties and moisture diffusion of vercors concrete. with the basic material parameters set, attention was paid to the behavior of the inner containment wall subject to ambient conditions and regular pressure tests. the conclusions can be summarized as follows. • a computationally efficient model was developed to study the influence of cyclic temperature and relative humidity on concrete creep. such an extensive study could not have been executed and analyzed using a full computational model of the containment. • the original version of the mps model is not suitable for simulating concrete subject to cyclic temperature. based on the data from the literature, the computed unrealistically high compliance was reduced by incorporating the parameter kt c set to 0.05 kt m. • the values of the most important material parameters were determined based on the provided experimental data. a simplified history of the ambient conditions suitable for the blind prediction of the containment was selected. the developed simplified model of the containment wall will serve for checking the plausibility of the structural response of the full computational model. acknowledgements the authors gratefully acknowledge financial support from the czech science foundation (ga čr), project number 19-20666s, and from the grant agency of the czech technical university in prague, project number sgs21/037/ohk1/1t/11. references [1] direction production ingeniere, edf. project maquette vercors, plan general, 2012. [2] štěpán krátký. benchmark vercors 2022 blind prediction of mechanical response of reinforced concrete containment, bc. thesis (in czech). czech technical university in prague, 2021. [3] direction production ingeniere, edf. specification of vercors benchmark 3, 2021. [4] z. p. bažant, l. j. najjar. nonlinear water diffusion in nonsaturated concrete. materials and structures 5:3–20, 1972. https://doi.org/10.1007/bf02479073. [5] z. p. bažant, a. p. hauggaard, s. baweja, f. j. ulm. microprestress solidification theory for concrete creep. i: aging and drying effects. journal of engineering mechanics 123:1188–1194, 1997. [6] z. bažant, s. baweja. creep and shrinkage prediction model for analysis and design of concrete structures: model b3. adam neville symposium: creep and shrinkage structural design effects 2000. [7] z. bažant, p. havlásek, m. jirásek. microprestress-solidification theory: modeling of size effect on drying creep. in n. bicanic, h. mang, g. meschke, r. de borst (eds.), computational modelling of concrete structures, pp. 749–758. crc press/balkema, eh leiden, the netherlands, 2014. 31 https://doi.org/10.1007/bf02479073 acta polytechnica ctu proceedings 34:26–31, 2022 1 introduction 2 materials and methods 2.1 laboratory experiments 2.2 material models 2.2.1 heat and moisture transport 2.2.2 structural analysis 2.3 computational models and calibration strategy 3 results and discussion 3.1 overall behavior 3.2 response to ambient conditions and pressure tests 4 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.35.0049 acta polytechnica ctu proceedings 35:49–53, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the long-pitch corrugation development in small radii curves jan valehrach∗, tomáš říha, jiří vendel, otto plášek brno university of technology, faculty of civil engineering, institute of railway structures and constructions, veveří 331/95, 602 00 brno, czech republic ∗ corresponding author: valehrach.j@fce.vutbr.cz abstract. the paper deals with the phenomenon of long-pitch corrugation in curves of the small radii and also the effects that accelerate or slow down its development. the paper presents data from sections of railway track measured for new rails, rails before and after grinding or replacement. the effect of train speeds running through the curves and the associated cant deficiency or excess were studied. the tendencies of long-pitch corrugation development are described, which can be used to estimate when maintenance action or rail replacement will be necessary. keywords: cant defeciency, cant excess, long-pitch corrugation, speed, track, train. 1. introduction various technologies to remove rail defects such as the long-pitch corrugation are present. the spectrum starts with the preventive grinding and spans all the way to a complete replacement of the rails. to understand which technology is for the track section or even for the specific curve at the time the most durable, a broad microgeometry data collection is necessary. based on the data history it is then possible to predict the further defect development and to lay down the maintenance guidelines in order to improve the possessions plan and to minimize negative effects on the railway operation. although there are procedures and methods to measure a microgeometry of rails, in most cases these applications are simplified for practical use in the track record, for compiling substantial amounts of data, or for the "rail reprofiling reception" [1]. the evaluation for the purposes of the development of the long-pitch corrugation then happens only based on line manager requests. for the purposes of measuring and monitoring the defects are categorized according to čsn en 13231 3 [1] into four wavelength ranges: • 10 – 30 mm, • 30 – 100 mm, • 100 – 300 mm, • 300 – 1000 mm. according to [2], wavelengths in the range 100–300 mm are considered as long-pith corrugation (corrugation primarily on the surface of inner rails in curves). the categorization mentioned above is not very suitable for the long-pitch corrugation monitoring, hence the length of the waves could be even shorter than 100 mm [3, 4], especially in the curves of small radii. following [5], based on the long-pitch corrugation creation theory, own measurements, and data comparison, it could be concluded, that for the long-pitch corrugation monitoring the range 30-100 mm and the range 100-300 mm may be merged, since wave amplitudes in both ranges usually reach similar or even same values. therefore, according to the above stated, for the rail long-pitch corrugation in particular the ranges 30-100 mm a 100-300 mm are to be monitored. 2. the long-pitch corrugation development monitoring, evaluation, and data quantification of the measured data were conducted using the rms and p2p calculations. the assessment originates from the 2006 standard čsn 13231-3:2006 [6], and not from the actual 2012 standard čsn 13231-3:2012 [1] due to the exclusion of the rms method. to be able to use the data measured before the acceptance of the 2012 standard, older standard was used. all the measurements were conducted on the railway line no. 326 (brno – česká třebová) between traffic point brno – maloměřice st. 6 and railway station adamov. own measurements took place in sections called hady (ha1.1 and ha1.2 to ha1.4), bílovice (bv1.1 and bv1.2) and babice (bs1.2), the track recording car data were used for the section adamov (ad1.1). once the calculated values rms and p2p are inserted into the graphs, the development speed gradient of the monitored magnitude could be obtained. examples of the graphs for measuring places in section hady are shown in figure 2 and figure 3. the line-up of the measuring places according to the gradient is shown in table 1. the steeper gradients are positioned lower. the consecutive order is also affected by the wavelength range. when the weighted cant deficiency (or cant excess 49 https://doi.org/10.14311/app.2022.35.0049 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en j. valehrach, t. říha, j. vendel, o. plášek acta polytechnica ctu proceedings figure 1. placement of the sections on the railway line no. 326 (brno – česká třebová). figure 2. rms progression in time for section hady. respectively) is calculated for each measuring place (curve) as shown in table 2, the rank of the values in comparison with the tab. 1 matches the wavelength range 100-300 mm. figure 3. p2p progression in time for section hady. 30-100 mm 100-300 mm bs1.2 bs1.2 bv1.2 ha1.1 bv1.1 bv1.2 ha1.1 ha1.2-4 ha1.2-4 bv1.1 table 1. measuring places line-up according to gradient. curve lweighted [mm] bs1.2 21.0 ha1.1 46.1 bv1.2 67.6 ha1.2-4 73.8 bv1.1 87.0 table 2. weighted cant deficiencies in curves for all recorded trains (not rounded). 3. development prediction 3.1. prediction based on the previous defect progression sections hady and bílovice using the gradients, the timeframes required for reaching the rms and p2p values were calculated for the purposes of the prediction of the future defect development. the timeframes respect the replacement works, or the rail grinding according to the particular place or section. the timeframe size differs significantly according to 50 vol. 35/2022 the long-pitch corrugation development in small radii curves section hady bílovice measuring place 1.1 1.2 1.3 1.4 1.2 number of years 0 – 24 40 – 63 0 – 24 40 – 63 0 – 24 40 – 63 0 – 24 40 – 63 from timeframe 30-100 rms 5.1 6.9 4.1 5.1 4.9 5.7 3.6 4.5 5.0 p2p 4.7 6.3 4.0 4.2 4.7 5.6 3.4 5.3 5.1 100-300 rms 22.0 5.6 28.3 37.0 21.0 17.4 49.7 14.2 4.7 p2p 23.0 4.9 26.1 30.3 23.3 14.9 47.7 13.7 3.9 table 3. weighted cant deficiencies in curves for all recorded trains (not rounded). figure 4. number of years needed to reach the wave defect progression of the same level as before the rails replacement. the wavelength range. whilst the wavelength range of 30-100 mm indicates the timeframe in years, the wavelength range of 100-300 mm reaches tens of years in timeframe estimation. the interval of 17 to 19 years could be associated with the essential modernization of the railway line at the end of the 1990s, where the sections are placed. on the other hand, to assume that the defect development lasts for two decades would not be right, since major maintenance efforts such as rail grinding took place a number of times. lastly, the considerably longer timeframe for the defect progression for the measuring place ha1.4 is not the result of the slow development of the defects, but the reason probably is the significantly worse state of the rails before their replacement in 2015. when calculating the timeframe required for reaching the defect progression of the same level as before rail grinding, the results are shorter. the reason is given by the lesser absolute values to be reached on one hand, but also by the initial state of rails on the other, hence the state of rails after grinding does not necessarily have to be equal to new rails. unfortunately, for the measuring places bílovice 1.1, babice 1.1, and 1.2 the initial state before the replacement of the rails in 2016 was not documented. section adamov the same principles of the development prediction were applied to the data from the track recording car in the section adamov. figure 5. number of years needed to reach the wave defect progression of the same level as before the rail grinding. the p2p from the track recording car is simplified just to values in 0.25 m steps, where all 4 values representing one meter of the rail are given the same value. these values are for the purposes of this paper signed as p2p modified (p2pm). furthermore, when all the values for the measuring place are summarized, the p2p modified sum is obtained (p2pms). this simplification regrettably brings the loss of information about the development progress, on the other hand, it helps to reduce local inaccuracies. when the p2pms values are drawn into the chart, the gradient could be assessed. the drawn values are referred to the initial state of the new rails, i.e. after the rail replacement. together with the known p2pms values before the rail replacement, it is then possible (by extrapolating the gradients) to search for the time, when the p2pms reaches the same value as it was before the rail replacement. this search is practically possible to be conducted on all the four wavelengths, however, the soundest ranges to be examined are the 30-100 mm and the 100-300 mm. to clarify: the wavelength range 1030 mm, which is embodying the non-periodical rail defects, could show locally randomized development progress. contrariwise, the wavelength range 3001000 mm represents more the defects of the track or even the defects of the whole substructure and do not reflect the maintenance done on the superstructure. the referential value of the developed defect for 51 j. valehrach, t. říha, j. vendel, o. plášek acta polytechnica ctu proceedings section hady measuring place 1.1 1.2 1.3 1.4 number of years 0 – 24 40 – 63 0 – 24 40 – 63 0 – 24 40 – 63 0 – 24 40 – 63 from timeframe 30-100 rms 2.2 3.0 2.3 2.8 2.4 2.8 2.4 3.0 p2p 2.2 3.0 2.2 2.4 2.4 2.9 2.4 3.8 100-300 rms 3.8 1.0 3.9 5.1 2.7 2.2 5.0 1.4 p2p 3.8 0.8 3.3 3.8 2.8 1.8 4.4 1.3 table 4. number of years needed to reach the wave defect progression of the same level as before the rails were ground. figure 6. wave defects development prediction based on the previous measurements done by the track recording car. the measuring place adamov 1.1 was taken from the measurement in 2016. maintenance works were carried out in 2018 (the values for the gradient obtention were from 2018, 2019, and 2020). the type of the regression curve (linear, exponential, . . . ) is based on the coefficient of determination so that its value is as close to 1.0 as possible. the p2pms value corresponding with the initial state before the maintenance works was for the measuring place adamov 1.1 found in 2035, which represents the progression of the defect for 17 years. it is interesting that this figure is comparable with the progression between 1999 (the newly modernized railway line put into service [7]) and the maintenance works in 2018. the development prediction, however, could not reflect some minor maintenance works, and also due to not very rich set of measured values, it is burdened with some uncertainty. moreover, the maintenance works in 2018 were of much lesser scale in comparison with the complex modernization in 1999. that is why the future timeframe still remains to be misty (see figure 6). 3.2. prediction based on the given limits the most suitable method for prediction is to use the gradients for calculation of rms and p2p into the values table 17 (čsn 13231-3:2006 [6]). figure 7. timeframe needed for the defect progression according to table 17 (čsn 13231-3:2006 [6]). the same applies to the standard sždc s2/4 [8] for the level 0.1 mm. figure 8. timeframe needed for the defect progression according to standard sždc s2/4 [8] for the level 0.1 mm. the longest timeframe for the defect progression or development appears to be in babice, although there were some negative gradients for the wavelength range 30-100 mm excluded from the assessment, and also the rms value for the wavelength range 100-300 mm is cropped. 52 vol. 35/2022 the long-pitch corrugation development in small radii curves 4. conclusions based on the measurements of the actual state there is the possibility to predict the speed of the further development of the long-pitch corrugation in curves of small radii. simultaneously, there is a possibility by monitoring weighted cant deficiencies (respectively cant excesses) to estimate in advance in which curves the progression of the defect will be faster. currently, a spectrum of measures exists to suppress the development of the long-pitch corrugation: starting with the methods aiming to prevent or slow down the defect progression and concluding with the measures designed to remove already developed defects. it is thus possible to use elastically enhanced superstructure or to grind off the defects extensively. that is why infrastructure managers face the difficult decision of which of the methods would be most suitable for their respective cases. once they had some appropriate tool to predict the defects progression, they can easily, based on the previously conducted measurements, decide for the actions generating lesser expenses or minimize the track possessions. performed measurements have also shown, that long-pitch corrugation monitoring via the p2p values is the most suitable way. assessments have additionally proven, that such monitoring should not be conducted without observation of other factors and parameters as well, hence the long-pitch corrugation is the result of the processes happening in an interconnected system infrastructure – vehicle – traffic. some parameters change over time very dynamically (rail wear, composition of trains), others are consistent (such as curve radii). long-term monitoring shows, that even in sections, appearing for the first sight as homogeneous, where unified construction and parameters take place, still some anomalies are present with different properties. these could be not only of the construction type (like the superstructure composition), but also even historical (previous actions not just to the infrastructure itself, but to the entire system infrastructure – vehicle – track). that amplifies the need for the selection of the appropriate parameters for the monitoring. regarding the current state, where the tracks are regularly diagnosed by the track recording car for the superstructure, there is no lack of high-quality data. in the case of the national line corridors, this means consistently three times during one year to obtain such data. naturally therefore there should be the drive to assess the data not only for the backward analysis but also for the defect development prediction. in the modern era, where powerful hardware and neural networks are present, data analysis is even easier. there is also the need for defect progression prediction on the designer side. the current state of the general knowledge of track designers is not overly broad: the usual conclusion is that the usage of curves of small radii is wrong, and it will eventually lead to long-pitch corrugation development. this is the space, where quality information about the parameters affecting the progression of long-pitch corrugation could be used most. track designers like to use guidelines and therefore some kind of a handbook, where the significant parameters, superstructure choices, and other crucial aspects would be summarised, would eventually lead to a significant decrease in the construction and maintenance costs and also in the number of possessions. as a result, society would obtain superior, safer, more dependable, and a neighbourly conscious railway track. references [1] čsn en 13231-3. železniční aplikace – kolej – přejímka prací – část 3: přejímka reprofilace kolejnic v koleji. praha: úřad pro technickou normalizaci, metrologii a státní zkušebnictví, třídící znak: 736374. 2012. [2] s. scholle. možnosti měřících vozů tudc železničního svršku pro detekci kolejnicových vad. příspěvek na workshopu: výhybky a výhybkové konstrukce. kolejnicové vady. brno, 2015. [3] c. esveld. modern railway track. 2nd ed. zaltbommel: mrt-productions, 2001. isbn 90-800-3243-3. [4] h. funke. brooušení kolejnic. praha: elka press, 1992. 122 s. [5] j. valehrach, p. guziur, t. říha, o. plášek. assessment of rail long-pitch corrugation. iop conference series: material science and engineering [online], 2017, roč. 236, č. 1, 8 s. issn: 1757-8981, https://doi.org/10.1088/1757-899x/236/1/012048. [6] čsn en 13231-3. železniční aplikace – kolej – přejímka prací – část 3: přejímka reprofilace kolejnic v koleji. praha: úřad pro technickou normalizaci, metrologii a státní zkušebnictví, třídící znak: 736374. 2006. [7] k. černý, r. jeschke, m. navrátil, et al. železniční trať brno – česká třebová na starých pohlednicích. polnička: tváře, 2019. isbn 9788090739476. [8] sždc s2/4. předpis pro zajišťování diagnostiky železničního svršku měřícími prostředky s kontinuálním záznamem. účinnost od 1. ledna 2020. správa železniční dopravní cesty, státní organizace. praha, 2020. 53 https://doi.org/10.1088/1757-899x/236/1/012048 acta polytechnica ctu proceedings 35:49–53, 2022 1 introduction 2 the long-pitch corrugation development 3 development prediction 3.1 prediction based on the previous defect progression 3.2 prediction based on the given limits 4 conclusions references https://doi.org/10.14311/app.2022.33.0591 acta polytechnica ctu proceedings 33:591–596, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague use of recycled concrete fines in cement and as aggregate simone stuerwald∗, ronny meglin, susanne kytzia, sabrina gilg open university, walton hall, milton keynes, keynes mk7 6aa, uk ∗ corresponding author: simone.stuerwald@ost.com abstract. the research project focused on investigating and optimizing the processing and use of recycled crushed sand 0/4 from concrete demolition waste, as an alternative raw material in the cement and concrete industry. crushed sand is produced during the processing of concrete demolition waste. the goal was to identify the optimum way of using the processed material along the entire process chain so that greenhouse gas emissions, waste volumes are reduced, and natural resources are conserved. different samples of laboratory and real crushed concrete fines were collected and examined in relation to various possible applications in accordance with the applicable standards. results highlight, that crushed concrete fines can be used in various applications in the concrete value-chain. however, for an optimal usage, additional processing is needed. keywords: cement, concrete, construction and demolition waste, lca, recycling. 1. introduction more than 10 years ago, a swiss standard was introduced that regulated the use of recycled concrete demolition waste as an aggregate for structural concrete [1]. nowadays the use of such aggregates is very common in europe and switzerland [2–4]. in the coming decades, the amount of construction waste in switzerland will increase especially concrete demolition waste from structures built in the 1950s to 1970s [5]. however, most of the processed construction waste is "down-cycled" into materials with subordinate technical requirements such as lean concrete or as road subbase material [6]. for this reason, various studies have recently been carried out to optimize the use of concrete demolition waste [7–9]. however, most studies focus on the complete concrete demolition waste. we focus on the crushed concrete fines because it has a significant impact on the possible use of the demolition material due to its less favourable properties [10]. we think, that the usage can be optimized, especially as it is assumed that recycling of concrete fine from demolition could further reduce co2 emissions in the concrete value chain [11–13]. we aim to optimize the usage of concrete demolition waste, especially the crushed concrete fines as a by-product. the economic efficiency of concrete recycling will be significantly improved if this crushed sand can be used profitably in cement and concrete production. we evaluate three economically viable uses of recycled crushed sand 0/4 according to sustainability criteria such as climate protection, resource conservation and waste avoidance. possible applications would be (i) use as a low-co2 substitute in cement kiln, (ii) use as an additive in cement mill or concrete production, or (iii) use as an aggregate in concrete production (figure 1). the aim is to identify the optimum place of use, the quality of the crushed sand required and the requirements for processing methods. in this way, crushed sand is to be optimally returned to the material cycle, natural raw materials are conserved, and co2 is saved. 2. methodology we estimate the potential for the use of crushed concrete fines by using various test methods and calculations to determine the properties of real and laboratory specimens. three samples of crushed sands (bs-l) were produced in the laboratory with different cement paste contents. another three samples (bs-r) were taken from recycling plants and a sample of sand from natural resources was used as a reference. based on chemical analyses of the samples, the use of ground crushed fines in clinker production is evaluated in simulations and the possible degree of substitution as a raw material substitute is estimated. subsequently, mortar samples produced with the ground crushed fines as an admixture are tested according to sn en 196. the results are then compared with the current standards and allow predictions on the possible use of the crushed sand as an additive and the possible amount to be added. in a further step, crushed fines samples are used as a substitute for primary sand in the production of concrete. the concretes produced are then subjected to extensive testing according to sn en 206. a comparison with the standards will allow us to estimate the maximum degree of substitution of the primary sand. finally, the results will be evaluated and discussed. 591 https://doi.org/10.14311/app.2022.33.0591 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en s. stuerwald, r. meglin, s. kytzia, s. gilg acta polytechnica ctu proceedings figure 1. in the process chain, raw materials are used at various processes: in the cement kiln (green), in the cement mill/concrete production (blue) and in concrete production (red). figure 2. thin section to determine the content of cement paste. 3. results 3.1. substitution of raw material in cement production recycled crushed fines contains sand and hardened cement paste and can be used as an alternative raw material in cement production [6, 10]. in particular, the cement paste portion contains the raw materials that can be reactivated in the cement kiln. crushed fines samples (bs-l) from the laboratory were specifically made with 40%, 55% and 70% cement paste. the cement paste content of real crushed fines (bsr) was determined on prepared thin sections by polarized light microscopy (figure 2). it was observed that the content of cement paste in the real crushed fines samples was 14% 15%, which is lower than the expected typical content of about 30% (figure 3). it is assumed that parts of the cement paste were lost in the demolition and separation process. material analysis by x-ray fluorescence spectroscopy (xrf) was performed on each of the real crushed fines samples (bs-r). the cao content was between 24 % and 29 % and the lsf (limestone saturation factor) was between 16.9 and 27.8. this reflects the low cement paste content, which is decisive for the maximum substitution of natural raw material. calculations showed that only small amounts of the raw material can be substituted. to raise the substitution rate, the cement paste of crushed sand must be separated completely. hardened cement paste by itself could replace more than 90% of the natural raw materials (figure 4). 592 vol. 33/2022 use of recycled concrete fines in cement and as aggregate figure 3. samples of crushed sand and their content of hardened cement paste. figure 4. maximum substitution rate depending on lime-saturation-factor (lsf). 3.2. scms in cement mill or concrete production to evaluate the possible use of crushed sand as an admixture, selected samples of crushed sand (bs-r1, bs-l-40, bs-l-55 and bs-l-70) were milled and mixed in a ratio of 40:60 % with cem i 52.5 r. these cements were then used to prepare cement paste for further testing of strength (according to en 196-1), setting time and room stability (both according to en 196-3). results of cem ii/b-ll were used as reference. the results of the tests (table 1) indicate, that a lower compressive strength compared to a cem ii/bll must be expected. since the strengths of cements with crushed fines do not differ significantly, it can be assumed that the crushed fine is inert. the results of the setting time do not show a clear picture. while the cement with bs-r-1 crushed fine sets significantly faster than a cem ii/b-ll, the cement with bs-l-40 shows a somewhat slower setting time. the room stability test does not indicate any significant differences between the cements tested. the processing and preparation of the crushed sand as an additive is more complex than for the other two types of use since high demands are made on purity and homogeneity of the material. 3.2.1. substitution of sand in concrete production using the crushed fines as a substitute for natural sand is a good method to reduce the consumption of primary resources. however, sand is often segregated from demolition material because the fresh and hardened concrete properties may deteriorate. first, the particle size distribution (figure 5) and the properties of the crushed sands (table 2) were determined. in a test series on concrete mixtures, it was investigated whether samples of crushed sand are suitable as a substitute for sand and which concrete properties are derived. in each mixture, the entire aggregate fraction 0/4 was replaced by recycled crushed sand. 593 s. stuerwald, r. meglin, s. kytzia, s. gilg acta polytechnica ctu proceedings cem bs-l-70 cem bs-l-55 cem bs-l-40 cem bs-r-1 cem ii/b-ll water [%] 30.1 30 30 28 28 compressive strength [n/mm2] 2 d 21.3 21.6 20.9 20.8 25 7 d 30.8 34.9 32.4 33.3 37 28 d 36.9 41.4 39.4 39.2 47 flexural strength [n/mm2] 2 d 4.8 4.5 4.4 4.9 n/a 7 d 5.7 5.9 6 6.3 n/a 28 d 6.2 6.3 5.9 6.3 n/a setting time [min] begin 157 187 216 127 200 end 204 231 256 207 240 volume stability [mm] 0.5 1 1 1.5 1 table 1. results of cement paste tests compared to cem ii b/ll. figure 5. particle size distribution of crushed sands. a mixture with natural sand was used as a reference. the different water absorption of the crushed sand samples was considered, so that a constant effective w/c ratio of 0.6 should be achieved. the compressive strength of concrete with recycled crushed sand was only reduced by about 10-15%. however, this reduction may also be related to differences in the effective water content and need not be related to the mechanical properties of the crushed sand. crushed sand with high content of cement paste (bs-l-70) show a young’s modulus that is up to 17 % lower, while a mixture with low content of cement paste shows a young’s modulus that is about 6 % lower compared to the reference. natural sand can be completely replaced by recycled crushed sand. if the cement paste is separated during processing, the remaining sand can be used to produce concrete, which is technically equivalent to concrete made from primary material. 4. discussion the results show that the crushed fines can, with certain restrictions, be used for the investigated application possibilities in the value chain of concrete production. the results of the mortar tests (section 3.2.) indicate that a complete substitution of limestone powder is possible without any significant loss of quality. the use as a substitute for natural aggregates is also possible according to the results in 3.3. to do this, however, further concrete technology measures must be taken to prevent any significant loss of quality. the use as a raw material substitute in cement production (3.1.), however, does not seem to be possible without further processing steps. this is due to the low cement content found in the real samples, which leads to a low lime saturation factor and thus prevent a high degree of substitution. an aspect not presented here in detail is the co2uptake of the concrete waste. the cement paste absorbs co2 from the air through natural carbonation 594 vol. 33/2022 use of recycled concrete fines in cement and as aggregate bs-r-1 bs-l-40 bs-l-70 ratio of cement paste [%] 14 40 70 dry density [kg/m3] 2332 2175 1940 water absorption wa24 [%] 6.2 9.7 12.4 table 2. properties of the crushed sands. reference bs-l-70 bs-l-40 bs-r-1 cement [kg/m3] 300 300 300 300 water [kg/m3] 180 180 180 180 w/z (eff.) [−] 0.6 0.6 0.6 0.6 aggregate [kg/m3] 1887 1605 1678 1738 density [kg/m3] 2367 2085 2158 2218 air content [vol. %] 1.5% 1.5% 1.5% 1.5% table 3. concrete mix and properties of fresh concrete mixtures. reference bs-l-70 bs-l-40 bs-r-1 density concrete [kg/m3] 2400 2310 2300 2320 strength 28d [n/mm2] 40.5 36.0 37.0 34.5 e-modulus [n/mm2] 38ă419 32ă883 31ă867 35ă976 table 4. properties of hardened concrete mixtures. and binds this in the formation of calcium carbonate (caco3). this effect has already been investigated by various studies, although there is still no generally valid assumption about the exact maximum possible degree of carbonation [14–16]. carbonation occurs particularly when the cement stone is ground and therefore has a very large surface. the co2 uptake is particularly useful in the production of clinker, as the co2 emissions of the rotary kiln can be kept in the "cycle" by means of forced carbonation. in this way, we could achieve a certain degree of co2 neutrality in clinker production. based on these results, we would like to suggest the following approaches: • the greatest possible separation and recovery of the cement paste when processing concrete demolition. when developing processing methods, the aim should be to separate the cement stone as far as possible and to hold it back as a separate fraction. this gives you (i) a high-quality and climateneutral secondary raw material for cement production and (ii) a high-quality rc aggregate for concrete production, which enables the further optimization of recipes to produce rc concretes. • as far as possible carbonation of the cement paste in the entire process chain of the preparation of the cement paste in the concrete demolition (i.e. the preparation and storage in the various process stages). we assume that concrete demolition in switzerland has a significant potential for rebinding co2, since it is mostly concrete that was produced with portland cement and only absorbed limited amounts of co2 during use. in the development of processing methods, the aim should be to increase the carbonation of the cement paste in the concrete demolition. this is achieved above all by increasing the surface area and designing the storage conditions in a targeted manner. this allows co2 to be extracted from the air and stored in the building material. this corresponds to the "carbon capture and storage" approach that is much discussed today, with which one would like to reduce net co2 emissions by 2050. 5. conclusions the demolition of concrete is not yet considered as a problem and the potential of its recycling is not yet sufficiently recognised. this will change in the coming years or decades, because: (i) the amount of concrete demolition will increase once the concrete structures from the second half of the last century have reached the end of life. therefore, the proportion of mixed demolition in mineral construction waste will decrease in the future and the proportion of concrete demolition will increase accordingly. (ii) the possible uses of rc concretes in structural engineering will increase and with it the demands on the building material. today, qualitative deficiencies in the concrete granulate are compensated by adjustments in the concrete mix design (more cement, more admixtures). this leads to higher costs and environmental pollution. in the future, attempts will be made to improve the quality of concrete demolition. (iii) the pressure on the cement industry to make its contribution to climate protection is increasing (e.g. 595 s. stuerwald, r. meglin, s. kytzia, s. gilg acta polytechnica ctu proceedings through higher co2 fees and measures taken by public clients). many representatives of this industry see a promising approach to solving this problem by rebinding co2 in concrete. for this, the greatest potential lies in the rebinding capacity in the crushed sand of the concrete demolition. for these reasons, efforts to develop new processes for the separation and processing of c&dw will increase significantly in the coming years and further investigations for the usage of c&dw will be necessary. the here described research project improves the decision-making basis for the development of new technologies and possibilities to use a material which, until now, is down-cycled or landfilled. acknowledgements supported by the federal office for the environment foen (utf 591.03.19) and the gebert rüf foundation (grs-049/18). references [1] sia. recyclingbeton, sia 2030, 2010. zürich. 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[16] w. ashraf. carbonation of cement-based materials: challenges and opportunities. construction and building materials 120:558-70, 2016. https: //doi.org/10.1016/j.conbuildmat.2016.05.080. 596 https://doi.org/10.1016/j.resconrec.2018.04.016 https://doi.org/10.1111/jiec.12739 https://doi.org/10.1016/s1672-2515(07)60241-3 https://doi.org/10.1016/j.cogsc.2018.02.010 https://doi.org/10.1111/jiec.12262 https://doi.org/10.1016/j.wasman.2018.01.028 https://doi.org/10.1016/j.cemconres.2019.105819 https://doi.org/10.1073/pnas.1913867116 https://doi.org/10.1016/j.rser.2019.109495 https://doi.org/10.1007/s41024-016-0010-9 https://doi.org/10.1016/j.conbuildmat.2016.05.080 https://doi.org/10.14311/app.2022.33.0431 acta polytechnica ctu proceedings 33:431–436, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague structural flexural strengthening through material bonded to the concrete substrate victor hugo dalosto de oliveiraa, rodolfo alves carvalhoa, ∗, rodolfo rabelo nevesb, luana ferreira borgesa, paulo chaves de rezende martinsa, marcos honorato de oliveiraa, joão paulo de almeida siqueiraa, rafael de almeida sobrala, sidney luan teixeira de nazaréc, carlos henrique de moura cunhac a department of civil and environmental engineering, university of brasília, sg-12, campus darcy ribeiro, asa norte, brasília, df, 70910-990, brazil b department of materials engineering and construction, federal university of minas gerais, av. antônio carlos, 6627, pampulha, belo horizonte, mg, 31270-901, brazil c school of exact, architecture and environment, catholic university of brasília, qs 07 lote 01, taguatinga, brasília, df, 71966-700, brazil ∗ corresponding author: rodolfoalvescarvalho@yahoo.com.br abstract. this article proposes an alternative method for the structural design of reinforced concrete elements strengthened in bending by metallic plates or fiber-reinforced polymer (frp) bonded to the concrete substrate. it is proposed a new calculation procedure for the strengthening using thin adhered material bonded to the element surface that dispenses the iterative process generally used in the design. the proposed routine is validated by comparison with other methods. a practical example is also presented, applying the procedure to an element of a building where a load change was foreseen. as result, it was verified that the proposed procedure provides values similar to the trial-and-error method used in the frp strengthening design. results are also coherent with other methods available in the literature for metallic plates. therefore, since this routine obtains similar values without using an iterative method, its applicability in the design becomes advantageous. keywords: attached materials, calculation guide, structural strengthening. 1. introduction when constructions do not fit the required capacity for correct use, rehabilitation is necessary. the reasons for rehabilitation may be pathological manifestations, accidents, natural disasters or changes in the building use. to restore the necessary resistance of the building, actions such as recovery or structural strengthening may be necessary [1–3]. the primary function of a structural retrieval is to improve the resistant capacity of any structure. the strengthening can be performed by associating new materials to the original structure, forming a composite system that supports the applied loads. examples of these processes are the addition of steel bars to the original section, reinforced concrete jacketing, external prestressing tendons and elements adhered to the substrate, such as steel sheets and composite materials [4]. this study proposes a new calculation methodology for the flexural design of reinforced concrete elements associated with small thickness bonded materials to the surface, as external steel sheets or fiberreinforced polymer (frp). the main goal herein is to establish a calculation guide for structural strengthening design, which fits the requirements of the brazilian standard nbr 6118: 2014 [5]. it should be noted that only the flexural capacity is concerned in this paper. considerations related to anchor capacity or the possibility of debonding of the strengthening are not made. 2. literature review and numerical simulation this section presents some methods commonly used for the flexural strengthening design of reinforced concrete beams and slabs present in the literature. these methods consider the insertion of steel sheets and carbon fiber reinforced polymer sheets bonded to the surface. the calculation guide proposed in this paper follows the literature review. 2.1. calculation models for metal sheet strengthening there are several calculation methods in the literature for the flexural strengthening using bonded metal sheets and steel profiles. some of the most relevant are described below. 2.1.1. j. bresson model metal sheet dimensioning with bresson’s method is carried out by balancing the bending moments in relation to the most compressed fiber. materials are 431 https://doi.org/10.14311/app.2022.33.0431 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en v. h. d. d. oliveira, r. a. carvalho, r. r. neves et al. acta polytechnica ctu proceedings figure 1. stress diagram of a strengthened beam. source: adapted from silveira [6]. assumed to have a linear elastic behavior; concrete tensile strength is neglected and stresses are limited to their respective elastic admissible values [6]. thus, the strengthening plate area is obtained by equation 1: as = ! mdead + mlive " + (fc1 + fc2 ) b·y2 6 − (fs1 + fs2 ) as d fr · h (1) as shown in figure 1, design takes into account dead and live loads, referred, respectively, to the forces already acting in the section when the strengthening is carried out and after the rehabilitation. 2.1.2. campagnolo model in this model, strengthening design is performed using classical formula for bending, considering composite beams theory [1]. thus, the tensile stress calculation for the steel plate is obtained as follows in equation 2. fr = er ec · mu (h − xii ) ix (2) 2.1.3. cánovas model the design is performed on the basis that materials behave under their maximum allowable elastic stresses. stress distribution before and after the strengthening execution are considered. balance and design equations are obtained by superimposing the stress diagrams [1], as illustrated in figure 1. thus, the necessary area of additional strengthening, ar, is obtained by balancing external and internal moments as in equation 3: mu = (as fs + ar fr) (d − x′) (3) 2.1.4. ziraba and hussein model the authors proposed a method in which the strengthening design is done at the ultimate limit state, following the specifications of aci 318 [1, 7]. stress resultants and the strengthening area are obtained from the balancing of moments in the section taken at the point of the compressive stress resultant in concrete, as reproduced in figure 2, resulting in an iterative process, as shown in equation 4. msd = rs # d − y 2 $ + rr # hr − y 2 $ (4) 2.2. calculation models for carbon fiber reinforced polymer design of strengthening systems composed of carbon fibers, according to machado [2] and the american code aci 440 [8], may be done through a trialand-error procedure, that gives the required area of strengthening at the ultimate limit state. according to these methods, a position for the neutral axis (x) is initially arbitrated. then, deformations (ε) and stresses (f ) of materials are calculated using compatibility and equilibrium equations. it is then verified if the resulting forces are balanced (equation 5), and if the resistant bending moment gotten is greater than the external moment applied (equation 6). if any of the conditions is not satisfied, a new neutral axis depth is chosen, and the procedure is repeated. nsd = rs + r′s + rr + rc (5) msd = rs # d − yx 2 $ + r′s # yx 2 − d′ $ + rr # hr − yx 2 $ (6) 2.3. proposed calculation model this section contains the method proposed by the authors. it is a calculation guide used for structural strengthening design, addressing the association of thin-thickness materials adhered to the surface. 2.3.1. calculation premise the following hypotheses are adopted: • the design is carried out at the ultimate limit state, disregarding the concrete tensile strength. • plane sections remain plane after loading, and the deformations are proportional to their distance to the neutral axis (bernoulli criterion). 432 vol. 33/2022 structural strengthening through bonded material figure 2. stress and strain diagram for a section in the strengthened condition. • perfect adhesion exists between the concrete, the steel reinforcement, and the thin adhered material (navier criterion). • the cross-sectional failure is characterized when the specific deformation of any material reaches its limit (ultimate) value. figure 2 illustrates the behavior of a rectangular cross-section of a strengthened concrete element, submitted to simple normal flexure. the rectangular stress diagram presented in section 17.2.2 of the brazilian standard nbr 6118: 2014 [5] is adopted. the compression stresses in concrete are distributed in a rectangle of depth y, with peak stress equal to α · fcd, according to figure 2. the value of y is defined as (equation 7): y = λ · x (7) in which, λ is given by equation (8) for fck ≤ 50 mpa, and λ is given by equation 9 for fck > 50 mpa: y = 0.8 (8) y = 0.8 − (fck − 50) % 400 (9) the α parameter is defined according to the concrete class as follows: for concrete classes up to c50: α = 0.85 (10) for concrete classes from c50 up to c90: α = 0.85 & 1 − (fck − 50) % 200 ' (11) 2.3.2. longitudinal strains considering bernoulli’s hypothesis, the deformation of any point within the section is proportional to its distance to the neutral axis, given as a function only of its location (y) and the curvature radius (ρ). knowing that the strengthening material bonded to the element has a minimal thickness, about a few millimeters, it can be considered that its specific longitudinal deformation (εr ) is equal to that of the concrete fiber substrate to which it is glued. thus, the deformation of the material, taken as constant along its thickness, can be obtained by equation 12: 1 ρ = εc x = εs (d − x) = εrf (h − x) = ε ′s (d − x) (12) when the applied loads, including the dead load, are not eliminated before the strengthening application, the material added to the section is not stressed by all these loads. in this way, the original structure is subjected to pre-existing stresses and has already suffered deformations before strengthening. under these conditions, it is necessary to know these stresses to determine the actual stress acting on the strengthening. thus, strengthening deformations can be obtained using equation 13: εr = (εrf − εri ) ≤ εu (13) in which εrf is the specific deformation of the concrete substrate when all loads are applied, and εri is the specific deformation in the section at the moment of strengthening application. 2.3.3. failure criteria analyzing the failure criteria, the ultimate concrete deformations are defined in section 8.2.10.1 of the brazilian standard [5], by εc2 and εcu, as shown in equations 14 to 17: for concretes classes up to c50: εc2 = 0.2 % (14) εcu = 0.35 % (15) for concretes classes from c50 up to c90: εc2 = 0.2 % + 0.0085 % (fck − 50) 0.53 (16) εcu = 0.35 % + 3.5 % & (90 − fck) % 100 '4 (17) considering hooke’s law valid up to yield stress (fy ) of any steel element, it is assumed that the 433 v. h. d. d. oliveira, r. a. carvalho, r. r. neves et al. acta polytechnica ctu proceedings figure 3. strengthened beam scheme. source: adapted from machado [2]. stresses are proportional to the deformations. in this way, the tensile resultant in the pre-existent steel reinforcement (rs) and in the strengthening added to the element (rr) can be done by equations equation 18 and 19. rs = as · es · εs ≤ as · fsy (18) rr = ar · es · εs ≤ ar · fry (19) it is important to note that, as the carbon fiber has high tensile strength, linear elastic behavior is expected for this type of material until failure of the critical section. finally, to prevent excessive plastic deformation in critical sections, the maximum allowable elongations are limited to = 1%. 2.3.4. ductility criteria in reinforced concrete structures, proper measures must be taken to ensure adequate ductility of all structural elements up to the ultimate limit state. according to aci 318-14 [7], if the tensile stress in the reinforcement is greater than εs = 0,5%, the section is characterized as ductile, presenting warnings about imminent rupture, such as the presence of large deformations and cracks. if the steel deformation is less than this value, a fragile rupture condition is expected, with few warnings of an imminent collapse. section 14.6.4.3 of the brazilian standard nbr 6118: 2014 [5], in order to guarantee a ductile behavior for the slabs and beams, establishes that the depth of the neutral axis at the ultimate limit state must be limited to the following values: equation 20 for ≤ 50 mpa, and equation 21 for 50 mpa < fck ≤ 90 mpa. x ≤ 0.45 · d (20) x ≤ 0.35 · d (21) 2.3.5. strengthening equation the equilibriums equation of strengthening supposes that dimensions of the cross-section, the reinforcement position, and the material properties are known. the procedure described below must consider the balance of the section, the compatibility of deformations, and its failure mode. it was considered that the internal resisting bending moment must be at least equal to the external moment applied to the section. in this way, the equilibrium equation of moments can be carried out at the contact surface between concrete and the newly added strengthening. one of the variables is eliminated with this procedure. therefore, the equilibrium of moments in a strengthened section is: ( m = msd msd = rc (h − x′) + rs (h − d) + r ′s (h − d ′) msd = α fcd bw y # h − y 2 $ + as fs (h − d) + a ′s f ′ s (h − d ′) (22) knowing that the compressed concrete block is given as a function of equation 7, the position of the neutral axis can be obtained by as the positive root of equation 23, derived from equation 22 by algebraic manipulation: ! −0.5 α fcd bw λ2 " x2 + (α fcd bw λ h) x + [a ′s f ′ s (h − d ′) − as fs (h − d) − msd] = 0 (23) with the neutral axis position, equations 12 and 13 are used to identify the domain of ultimate limit state of the behavior of the section, and to determine the specific deformations along its height. the sectional stresses distribution from existing load action during the strengthening can be considered, taking 434 vol. 33/2022 structural strengthening through bonded material variable iterative process [2] suggested model(authors, 2021) εri 0.228 ! 0.272 ! εrf 7.000 ! 7.324 ! εr 6.772 ! 7.052 ! l · n 0.228 cm 22.31 cm fr 1544.0 mpa 1607.9 mpa strengthening area 0.588 cm2 0.563 cm2 table 1. calculated carbon fiber strengthened area. figure 4. representation of beam characteristics of the building. calculation model strengthening area difference suggested model 2.054 cm2 j. bresson 4.091 cm2 99.17 % campagnolo 3.684 cm2 79.36 % ziraba e hussein 2.905 cm2 41.43 % cánovas 2.286 cm2 11.30 % table 2. comparison of the calculated strengthening sheet area. into account the cracked section properties, if necessary. considering a cross-section submitted only to flexure forces, the resultant strengthening force is determined as presented in equation 24. ( f = nsd nsd = rc + rs + r ′s + rr rr = nsd − α fcd bw λ x − as fs − a ′s f ′ s (24) once the resulting force and specific deformation have been determined, the strengthening area necessary to resist to the applied moment can be calculated using equation 19. finally, it is still necessary to assess whether the section mode of failure satisfies the ductility criteria set out in section 2.3.4. it is worth to mention that the proposed equation 24 also allows the direct application of the environmental reduction factors, as exposed in aci 440 [8], to the strengthening resultant. 3. results and discussions a comparison between the design methods mentioned and the calculation model proposed in this paper was made. for this, some examples available in the literature and a study case were taken for. analyzing example 4.1 of the manual by machado [2], a concrete beam was supposed to be strengthened through the association of carbon fiber sheets. all necessary data about the beam, the applied loads, and the material properties are shown in figure 3. the proposed calculation routine was compared with the iterative procedure used in machado [2]. the design was carried out using the criteria of the brazilian standard nbr 6118: 2014 [5], adopting the same parameters for the materials. the comparative results are shown in table 1. the actual strengthening provided had an effective area of 0.594 cm2, composed of two carbon fiber layers, 0.165 mm thick and 18 cm width. it appears that the recommended procedure provides a value very close to the trial-and-error process for necessary carbon fiber strengthening area (less than 4.25%). this difference may be due to the tolerance criteria applied in the iterative process [2]. 435 v. h. d. d. oliveira, r. a. carvalho, r. r. neves et al. acta polytechnica ctu proceedings the second example is the case of a building, where the live load was foreseen to be increased. the additional strengthening was conceived through the association of metal sheets. figure 4 shows the characteristics of one of the beams that needed to be strengthened. table 2 shows the section sizing of the steel plate obtained from the different methods. it is essential to highlight that, although the proposed method presents similarities with that developed by ziraba and hussein both consider the strengthening design at the ultimate limit state, their results are quite different. this fact is due, not only, but mainly because of criteria that govern aci 318 [7] is different from those of nbr 6118 [5] and because of the consideration of pre-existing deformations considered in the proposed calculation. 4. conclusions this paper shows that it is possible to design fiberreinforced polymer strengthening using a different approach from the trial and error method, frequently used in technical design. the proposed method leads to encouraging results with reasonable accuracy if compared with the classical ones. no iterative calculations are needed. the method provides the depth of the neutral axis without evaluating the stress resultant of the internal reinforcement. consequently, it is possible to consider its application for other thin materials bonded to the substrate, such as metal sheets. there was a distinct difference in the results obtained with other literature models for sheet metal design. the authors attribute these divergences mainly to the difference between some calculation hypotheses adopted, such as the design based on ultimate limit state instead of the allowable stress design methods, where material resistances are limited to its limit of elasticity. the proposed method also allows the usage of the load and resistance factor design, the consideration of pre-existing stresses, and the application of environmental reduction factors directly to the strengthening, which are not possible in some other methods. thus, the numerical analysis shown that the proposed method presents good efficiency. more exhaustive laboratory tests, experimental and/or numerical, may be foreseen to improve and calibrate the parameters of the method proposed herein. acknowledgements this study was financed in part by the coordenação de aperfeiçoamento de pessoal de nível superior brasil (capes). references [1] v. c. m. d. souza, t. ripper. patologia, recuperação e reforço de estruturas de concreto, são paulo: pini, 1998. [2] a. p. machado. manual de reforço das estruturas de concreto armado com fibras de carbono, são paulo: viapol, 2015. [3] p. k. mehta, p. j. m. monteiro. concreto: microestrutura, propriedades e materiais 2, são paulo: ibracon, 2014. http://www.viapol.com.br/media/ 97576/manual-fibra-de-carbono.pdf. [4] a. p. a. reis. reforço de vigas de concreto armado por meio de barras de aço adicionais ou chapas de aço e argamassa de alto desempenho dissertação de mestrado universidade de são paulo, são carlos, 1998. [5] associação brasileira de normas técnicas. abnt nbr 6118: projeto de estruturas de concreto procedimento, rio de janeiro: abnt, 2004. [6] s. s. silveira. dimensionamento de vigas de concreto armado reforçadas com chapas coladas com resina epóxi dissertação de mestrado universidade federal fluminense, niterói, 1997. [7] american concrete institute 2014 aci 318-14: building code requirements for structural concrete, farmington hills: aci, 2014. [8] american concrete institute 2007 aci 440.2r-08: guide for the design and construction of externally bonded frp systems for strengthening concrete structures, farmington hills: aci, 2007. 436 http://www.viapol.com.br/media/97576/manual-fibra-de-carbono.pdf https://doi.org/10.14311/app.2022.33.0357 acta polytechnica ctu proceedings 33:357–362, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague high-performance concrete containing waste vitrified tiles diana mariakováa, b, ∗, kristina fořtováa, zuzana jirkalováa, b, tereza pavlůa, b, petr hájeka, b a university centre for energy efficient buildings of czech technical university in prague, třinecká 1024, 273 43 buštěhrad, czech republic b department of architectural engineering, faculty of civil engineering, czech technical university in prague, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: diana.mariakova@cvut.cz abstract. this contribution deals with the possibility to use waste vitrified tiles (vt) in high-performance concrete. silica powder normally used in high-performance concrete is replaced by waste vitrified tiles in 25, 50, 75 and 100 %. the basic experiments are tested, such as water absorption, flexural strength and compressive strength. the durability is examined, freeze-thaw resistance is tested in 25, 50, 75 and 100 freeze cycles. the results are evaluated and compared with the reference high-performance concrete, containing silica powder. the recipe of concrete is optimized based on these results. keywords: durability, high-performance concrete, waste vitrified tiles. 1. introduction the production of construction and demolition waste (cdw) raised almost up to 42 % in 2019 in czech republic. from 2017, the numbers are similarly high, ranging around 40 % [1]. however, in recent years there appears increasing tendency of recycling and reusing cdw. there are studies dealing with the possibilities of the cdw utilization and management of processing it [2–4]. building materials create almost a half of worldwide produced materials. due to the rapid industrial development the problems are appearing all over the world such as decreasing sources of non-renewable resources and environmental impact connected with the extraction, manufacturing and transporting these materials. at the same time, it is necessary not to omit the associated processes, which release exhaust gases such as carbon dioxide, or processes associated with the end of life of these raw materials and materials produced from them [5]. the problem of the necessity of landfilling these used materials is connected as well. sustainable development and the entire global problem with environmental impacts is related with the volume of produced concrete and the buildings made from this material [6]. replacing of primary raw materials needed to create concrete mixtures is convenient because of saving the primary sources and the energy resources required to obtain these raw materials. concrete is the most used produced material in the world. the high production is the reason of the studies searching for alternative components such as natural aggregates substitutes or cement replacement. there are studies dealing with the possibility to replace natural aggregate in concrete, due to the decreasing applicable sources of this material. the standards require the fulfillment of strict criteria for the selection of suitable aggregates for use in concrete, especially for road aggregates. the studies are examined the utilization of bricks, glass or used concrete as recycled aggregate in concrete [7–11]. the main problem is the possibility of an undesired alkali-silica reaction which needs to be discussed and evaluated [12]. the possibility of replacing (fully or partially) cement or cementitious materials are examined as well. researches are testing the appropriate amount of replacement or the convenient grain size[13–15]. verification of the selected material properties is essential for proper interaction of the components in concrete mixture. basic chemical composition, pozzolanic properties or toxicity [16] should be observed to valorize pertinency of future use. this paper deals with the use of waste vitrified tiles in concrete as a partial of full replacement or silica powder (cement replacement). vitrified tiles were selected as a waste to recycle with pozzolanic properties. the aim of this research is to verify basic properties and find the suitable amount of the replacement to preserve the special properties of high-performance concrete, such as low water absorption or high durability of the material. this paper describes behavior of concrete mixtures containing waste vitrified tiles compared with the reference high-performance concrete in the view of durability and mechanical properties, such as flexural and compressive strength. the testing methods are described and explained, the results are measured, 357 https://doi.org/10.14311/app.2022.33.0357 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en d. mariaková, k.fořtová, z. jirkalová et al. acta polytechnica ctu proceedings calculated and evaluated. based on the results, the recipe of concrete mixture is optimized and adapted for future use in construction industry. 2. materials and methods the mixtures were designed based on the original hpc recipe developed in department of architectural engineering [17]. silica powder was replaced by powder from waste vitrified tiles which is a ceramic material composed of feldspar and clay, almost without any pores. the material is produced under high pressure (350 − 450 kg / cm) and fired at temperatures above 1200 degrees celsius. the replaced part of silica powder was 25 %, 50 %, 75 % and 100 % by weight. original hpc recipe was used as the reference sample, needed during the evaluation and data comparison in the next chapters. 2.1. materials materials according the basic recipe were chosen from local sources and suppliers for preparation 5 types of high-performance concrete mixtures. these materials are: portland cement cem i 42.5 r (cem), natural aggregate in two fractions (na fine and coarse aggregate), silica powder (sp), microsilica, superplasticizers and water. silica powder was replaced by the waste vitrified tiles (vt) in the mixtures, the replacement is described in table 1. amount of cement, natural aggregate, microsilica and water remains same in the mixtures, the variable segment is the amount of used waste vitrified tiles as a silica powder replacement. more detailed information about the properties of used materials are not crucial for the experiments and the results of this work. based on the properties of vitrified tiles, there was an assumption of positive impact on the results. the pozzolanic properties of used waste material can have a positive effect on the resulting strength of concrete samples. this work aims to find the ideal percentage part to replace the silica powder. 2.2. methods basic experiments were made to analyse properties of high-performance concrete containing waste vitrified tiles. due to the vitrified tiles characterization water absorption, compressive strength and flexural strength were examined. the durability was verified as well in 25, 50, 75 and 100 freeze cycles. 2.2.1. water absorption another measured parameter was the absorption of the samples. water absorption of concrete samples is a necessary parameter measured in construction due to its use in structures exposed to external influences, such as water. after 28-days concrete hardening, the samples were weighed, stored in water, then reweighed and the water absorption rate was evaluated based on the measured parameters. 2.2.2. compressive strength there were two types of tested samples. first type was cube 150 × 150 × 150 mm. the tested surface ac was 150 × 150 mm. second type was the fragments with the tested surface ac 40 × 40 mm. fragments from the samples after flexural strength testing were used. in this contribution, the destructive method of compressive strength test was used. the procedure was followed the czech standards čsn en 12350-1, čsn en 12390-1, čsn en 12390-2, čsn en 125041 and čsn en 12390-4 [18–22]. the samples were tested under the speed 2400 n/s, peak 20 kn and the test is completed when the sample breaks, after that the maximum measured force is recorded. five samples from each mixture were tested and the average value was then calculated from the results. whole method is described in čsn en 12390-3 [23]. 2.2.3. flexural strength the tested samples were specimens with dimension 40 × 40 × 160 mm according to čsn en 12390-5 [24]. the flexural strength is defined in čsn en 1015-11 [25] by three-point press. the samples were tested under the speed 50 n/s, peak 2 kn and the test is completed when the sample breaks. three samples from each mixture were tested and the average value is calculated from the highest measured force. this work is dealing with three-point bending, it is possible to use another method four-point bending. 2.2.4. freeze-thaw resistance the tested samples were 40 × 40 × 160 mm. this test was made according to čsn 73 1322 [25]. the samples were cooling by water for four hours to -20 ◦c and then warming up for two hours to 20 ◦c. preparation before freeze thaw resistance test includes data recording. the samples were tested in 25, 50, 75 and 100 cycles, three specimens from each concrete mixture. the compressive strength and the flexural strength are measured after freeze-thaw cycles and effect of the cycles (25, 50, 75, 100) is examined and evaluated. 3. results 3.1. water absorption the water absorption was tested on the cube samples 50×50×50 mm on three specimens from each concrete mixture. the results of water absorption test are shown in the table below (table 2). due to similar character of the concrete mixtures it was expected the similar value of water absorption. all the mixtures values are ranging around 3 %, more detailed results are in table 2. 3.2. compressive strength compressive strength was tested on two types of samples. first type was cube specimen 150 × 150 × 150 358 vol. 33/2022 concrete containing waste vitrified tiles component (g) cem na sp ms h2o vt ref 650 1200 240 175 180 0 vt25 650 1200 180 175 180 60 vt50 650 1200 120 175 180 120 vt75 650 1200 60 175 180 180 vt100 650 1200 0 175 180 240 table 1. recipe of the tested hpc mixtures. mixture weight of wet sample (g) weight of dry sample (g) water absorption ref 294.9 285.9 3.124% vt 25 300.4 292.7 2.642% vt 50 293.8 285.7 2.823% vt 75 292.8 283.7 3.220% vt 100 294.7 285.5 3.199% table 2. water absorption of tested high-performance concrete samples. mixture strength after 28 days (mpa) σ percentage relation to ref ref 119.32 4.75 100.00% vt 25 121.64 3.51 101.94% vt 50 112.89 3.47 94.61% vt 75 122.42 3.72 102.60% vt 100 116.34 6.19 97.50% table 3. compressive strength (cubes 150 × 150 × 150 mm). mixture strength after 28 days (mpa) σ percentage relation to ref ref 128.50 4.64 100.00% vt 25 122.96 5.42 95.70% vt 50 125.49 4.54 97.66% vt 75 131.29 3.66 102.17% vt 100 129.39 2.35 100.70% table 4. compressive strength (fragments of blocks 40 × 40 × 160 mm). mixture strength after 28 days (mpa) σ percentage relation to ref ref 12.31 0.59 100.00% vt 25 11.05 0.13 89.76% vt 50 12.60 0.28 102.24% vt 75 12.14 0.76 98.64% vt 100 12.54 0.34 101.84% table 5. flexural strength (blocks 40 × 40 × 160 mm). 359 d. mariaková, k.fořtová, z. jirkalová et al. acta polytechnica ctu proceedings mm, the tested surface was 150 × 150 mm. the results are shown in table 3. the values are raging 112.89 − 122.42 mpa, highest results were achieved by the mixture containing 75 % and 25 % of vitrified tiles. the reference samples hpc showed slightly lower values, the percentage relation to the ref samples is also shown in the table below (table 3). the lowest values have the concrete mixture containing 50% of waste vitrified tiles. however, the value is still high up to 94.61 % compared to the ref hpc sample. the second tested type of sample were fragments of blocks 40 × 40 × 160 mm. the used specimens were tested on the flexural strength before, the tested surface was 40 × 40 mm. summarized results are in table 4. the results were ranging 122.96 − 131.29 mpa and were similar as the results from the cube compressive strength testing. the best results were measured on mixture vt 75 in both cases (on cubes 102.6 % compared to the reference sample, on block fragments 102.17 % compared to the reference sample. 3.3. flexural strength flexural strength was tested on specimens 40 × 40 × 160 mm. the blocks were made in sets of three samples from each concrete mixture. the average values were calculated based on the maximum measured force and are shown in table 5. the highest values were measured on the mixtures containing 50 % and 100 % of waste vitrified tiles substitution. in the comparison with the reference samples, the lowest value has the mixture vt 25 which is observed in the most measured parameters (compressive strength on the fragments of blocks and water absorption). 3.4. freeze-thaw resistance the samples were tested in the sets of three samples. each sample was tested in the flexural strength and compressive strength. samples were exposed to 25, 50, 75 and 100 freeze-thaw cycles to confirm the ability of high-performance concrete to resist freeze-thaw cycles. the results validate that there is no negative effect to this quality caused by the addition of new material waste vitrified tiles and all the values are summarized in figure 1 and figure 2 below. the flexural strength results of reference sample (ref) showed increasing tendency depending on the increasing number of cycles. the values went from 11.54 mpa (strength after 28 days) up to 17.64 mpa (strength after 100 cycles). this tendency was observed on the samples from mixture vt100 and vt25. the increasing tendencies did not occur on samples vt50 and vt75. however, the samples did not show any cracks or signs of failure and after all the minimal measured strength is still 82 % (vt50) compared to the ref. on the other hand, the average values measured on mixture vt100 achieved 102 % compared to the ref. the compressive strength was measured on the fragments of samples 40 × 40 × 160 mm. the tested surface was 40 × 40 mm. in the compressive strength experiment, mixture vt75 showed the best results, where the strength achieved 105 % compared to ref. on the other hand, the lowest result was measured on vt100, where the strength stayed on 93% compared to ref. nevertheless, the result is satisfying in the relation with the other average values raging 95 − 98 % after 100 freeze-thaw cycles. the compressive strength after 25 cycles is ranging from 98 − 104 % compared to the ref, 92.5 − 105 % after 50 cycles and 91 − 98 % after 75 cycles. it is possible to say, that the results met the expectations. 4. conclusion four mixtures containing different amount (25 %, 50 %, 75 % and 100 %) of waste vitrified tiles as a substitutional material were designed and tested (mixtures vt25, vt50, vt75 and vt100). all the results were compared with the reference sample ref (high-performance concrete made from primary raw materials). the water absorption was experimentally tested, the absorption of mixtures with tiles was similar to the ref. the absorbency values of all tested samples were around 3 %. the flexural strength was tested and vt50 and vt100 showed the highest results. the measured force and calculated strength were higher than the values of the ref. similar numbers were evaluated on the samples after compressive strength experiment. however, the highest results were provided by vt25 and vt75 (cubes) and vt50 and vt100 (fragments of blocks). all these results were higher than results of ref. the freeze-thaw resistance was verified. the freezing cycles don’t have the impact to the results. the compressive strength of the samples made from the mixtures containing 25, 50, 75 and 100 % substitution of silica powder by waste vitrified tiles was raging 91 − 105 % compared to the reference sample. the flexural strength achieved 116 % compared to the reference sample after 50 freeze-thaw cycles, rest of the samples were ranging from 82 % up to mentioned 116 %. when omitted maximal and minimal, the average value was 99 % compared to the ref. based on the experiments and the results, it is possible to say that the proposed mixtures are suitable for further testing and verification of their properties, which so far confirm the possibility of their use as hpc. the similarity of the resulting properties of the tested mixtures containing waste tiles with the reference mixture is given mainly by the composition of used waste vitrified tiles. therefore, the resulting values correspond to the assumptions made at the beginning of this experiment and forms a solid basis for further research. 360 vol. 33/2022 concrete containing waste vitrified tiles figure 1. flexural strength values after freeze-thaw resistance experiment. figure 2. compressive strength values after freeze-thaw resistance experiment. acknowledgements this research was funded by ministry of education, youth and sports ltain19205 design of performance based concrete using sand reclaimed from construction & demolition wastes. this work has been supported by the grant agency of the czech technical university in prague, grant no. sgs19/093/ohk1/2t/1 use of waste materials in hpc in design applications and sgs21/096/ohk1/2t/11 influence of chemical properties on the use of waste materials in concrete applications. references [1] český statistický úřad, 2021. https://www.czso.cz/csu/czso/cesko-v-roc e-2019-vyprodukovalo-37-mil-tun-odpadu. 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[25] čsn 73 1322: determination of frost resistance of concrete, 1968. 362 https://doi.org/10.4028/www.scientific.net/msf.866.58 https://doi.org/10.1016/j.conbuildmat.2008.07.023 https://doi.org/10.1016/j.jksues.2020.04.013 https://doi.org/10.3390/su12010250 https://doi.org/10.28991/cej-2020-03091620 https://doi.org/10.2139/ssrn.3577292 https://doi.org/10.1016/j.matpr.2020.09.164 https://doi.org/10.3390/ma14020463 https://dspace.cvut.cz/handle/10467/54346 https://doi.org/10.14311/app.2022.33.0489 acta polytechnica ctu proceedings 33:489–496, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague prestressed ultra-high performance concrete (uhpc) beams for reusable structural systems: design and testing dario redaellia, ∗, jonathan moixa, alex muresana, b, jan brüttingb, corentin fivetb a institute of construction and environmental technologies, university of applied sciences western switzerland (hes-so), fribourg, switzerland b structural xploration lab, swiss federal institute of technology (epfl), fribourg, switzerland ∗ corresponding author: dario.redaelli@hefr.ch abstract. with the aim to reduce the environmental footprint of buildings, this paper presents an original structural system concept for two-way slabs in residential and office buildings. the proposed system extends the best practice in terms of modularity, versatility, demountability, reusability, and durability. a finite set of elements can be used to form the main load-bearing system of multiple successively constructed buildings having different and unpredicted layouts and static systems. ultra-high performance concrete (uhpc) was identified as one of the most promising materials for this application because of its high strength, extreme durability and the opportunities it opens for shape optimization and material consumption reduction. in the first part of this contribution, the main features of the new structural system and preliminary design assumptions for uhpc modules are briefly outlined. then, the authors present and discuss the results of an experimental campaign on prestressed uhpc beams that were tested in bending and shear, under service and ultimate load conditions. beams with transversal openings in the web were also tested to assess the influence of these openings on the cracking behavior and shear strength of the beams. experimental results provide useful information for the subsequent modeling of the structural system. thanks to the presence of fibers, transversal openings in the web only have a limited influence on the response of the beams, thus allowing horizontal technical shafts to pass through the structural thickness of the slab. keywords: building, reuse, slab, uhpfrc. 1. introduction today, the global resource consumption is 1.7 times higher than what the planet can sustain [1]. the construction industry alone is responsible for approximatively one third of material usage [2], energy consumption [3] and waste production [4] worldwide, with load-bearing systems (e.g. slabs) contributing the major share of these impacts. existing strategies to face these issues are mainly based on the use of construction materials with a low environmental impact (e.g. natural, locally available, or recycled materials) or on the design of material-efficient systems (e.g. weight minimization). these approaches, however, do not provide a long-term solution since they are not addressing in a decisive way the building’s end-of-life: costs and energy required to demolish, transport and transform materials to be recycled, and the inevitable degradation (down-cycling) of their properties. an alternative strategy to reduce environmental impacts [5, 6] consists in conceiving building’s components in order to allow their reuse with minimal interventions in successive buildings and for multiple building lifetimes. this strategy, even though not yet widely established from a technological point of view, has the potential to radically transform the construction industry towards a more sustainable circular economy. 2. a reusable structural system for two-way slabs the fundamental requirements that a building system must fulfil in order to ensure open-ended reusability are: durability, versatility, modularity, reversibility and adaptability [5]. a new type of structural system for two-way slabs, which complies with all the above requirements, has been conceived within the framework of a collaborative research project between the swiss federal institute of technology (epfl) and the university of applied sciences western switzerland (hes-so). the system (figure 1.a) consists of quadrilateral slab modules that can be interconnected in a plane to create slabs with different ground plans. each module (figure 1, right) is conceived as a grid of orthogonal and regularly spaced beams, creating an array of transversal openings that define possible locations both for columns (with the addition of a head capital 489 https://doi.org/10.14311/app.2022.33.0489 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en d. redaelli, j. moix, a. muresan et al. acta polytechnica ctu proceedings figure 1. schematic representation of the structural system (a) and of individual modules realized with gluelaminated timber (b) or uhpc (c). shear strength bending strength stiffness weight vrd [kn] mrd [knm] ei [knm2] [kg/m] 1 layer 11.5 8.8 719 8.7 2 layers 23.1 34.0 5750 17.3 table 1. geometric and mechanical properties of glue-laminated timber beams. figure 2. glue-laminated timber beams. to support the slab) and for vertical technical shafts. slab modules can also be stacked vertically to locally increase the static height of the system, thus allowing to tailor the strength and stiffness distribution of the slab to meet specific static requirements [7]. extremely versatile, the proposed structural system allows for multiple re-adaptations in successive buildings with unpredicted uses, geometries and support positions. for practical applications, the system was designed to resist vertical loads of residential and office buildings with spans varying between 4 and 12 meters (with multiple layers). a first implementation is presented here as a reference and was designed in gl32h glue-laminated timber. bolted steel plates fixed to the ends of the beams with epoxy glued steel rods were chosen as connections between modules (figure 1.b). the geometric and mechanical properties of timber beams are given in table 1 for one-layer and two-layer configurations (figure 2). each beam is made of eight 80 × 30 mm laths. due to the crossing of orthogonal beams, however, only half of the laths are continuous in each direction ((figure 2, left), which limits the structural performance of this solution. shear strength, bending strength and bending stiffness given in table 1 were calculated according to swiss standards [8] by assuming that only continuous laths contribute, which was confirmed by laboratory experimental tests. despite the lightness, the positive environmental properties and the good mechanical strength of timber, stronger and stiffer alternative solutions were searched. ultra-high performance fiber-reinforced concrete (uhpc) was identified as a promising alternative due to its very high strength, extreme durability and the opportunities it opens for shape optimization and material consumption reduction. the rest of the paper focuses on design, testing and preliminary modeling of the uhpc implementation. 490 vol. 33/2022 prestressed uhpc beams vr mr ei weight [kn] mrd [knm] [knm2] [kg/m] uls design values 46.5 36.2 6105 41.8expected average values 69.7 46.3 required average values 34.7 43.5 table 2. geometric and mechanical properties of uhpc beams. figure 3. uhpc beams. 3. conceptual and detailed design of the uhpc implementation with respect to timber, uhpc is heavier, more expensive and its production requires more energy and non-renewable materials. to provide a competitive solution with respect to timber, the following criteria were set as basis of uhpc modules design: • provide with a single uhpc layer (24 cm height) strength and stiffness equal or higher than two layers of glue-laminated timber (48 cm, see table 1) • optimize the geometry of the cross-section to reduce self-weight and material consumption • explore the potential to create transversal openings in the webs for horizontal technical shafts, which is not possible for glue-laminated timber beams detailed design of the beams was carried out according to swiss standard sia 2052 [9] assuming the mechanical properties of a relatively common tensilehardening uhpc mixture (class ua, [9]): compressive strength fu ck = 150 n/mm2, elastic limit in tension fu tek = 7.5 n/mm2, maximal tensile strength provided by the fibres fu tuk = 9.5 n/mm2 and modulus of elasticity eu = 55′000 n/mm2. table 2 gives the calculated uls design values of shear and bending strength, as well as the corresponding average values (that might be expected in case of testing). also given are the estimated average values of shear and bending strength required in case of testing so that the corresponding design values (uls) are at least equal to those of two layers of timber (see table 1 for comparison). uhpc beams are pre-tensioned with 12 mm diameter y1570c pre-stressing wires directly fixed to steel plates at the extremities of the beams (figure 1.c). since the position of a specific slab module within the structural systems is a priori unknown, each module must be able to resist both positive and negative moments. consequently, symmetrically prestress was designed, with one prestressing wire per flange. flanges were sized to resist the compressive force generated by bending moments at ultimate, while their shape was chosen to allow placing longitudinal and transversal prestressing wires with a minimal cover according to [9]. a web thickness of 30 mm (figure 3) was chosen to facilitate the casting of uhpc and to obtain a design value of the shear strength (vrd = 46.5 kn, table 2) significantly higher than the target design value corresponding to two layers of glue-laminated timber beams (vrd = 23.1 kn, table 1). thanks to this shear strength surplus, the creation of transversal openings in the web of uhpc beams can be explored. despite the presence of few experimental and modeling approaches in the literature [10–13], no recognized design approach exist to the authors’ knowledge for the shear strength of uhpc beams with transversal web openings. 4. experimental campaign on pre-stressed uhpc beams with or without web openings to acquire a better knowledge on the influence of web openings on the behavior of the studied system and to provide basic information for further analytical and numerical modeling, an experimental campaign on prestressed uhpc beams with and without web openings was carried out. 4.1. description of the test series four beams were fabricated for testing (figure 2). beams s1 and b1 have no web openings, while beams s2 and s3 have openings with different geometries. beams s1, s2 and s3 were designed to fail in shear and were tested in three-point bending with 1.5 m span. these beams were fabricated with a length of 1.5 times the test span, two test regions in the outer thirds and a strong region in the central part. by changing the position of load and supports (black and red configurations, figure 2), it was possible to carry out two tests per beam, named "side a" and "side b" hereafter. to avoid any risk of bending failure, the 491 d. redaelli, j. moix, a. muresan et al. acta polytechnica ctu proceedings figure 4. geometry of the four tested beams. beam / batch compressive strength fu cm* [n/mm2] elastic modulus eu cm* [n/mm2] elastic limit in tension fu tem** [n/mm2] tensile strength fu tum** [n/mm2] hardening strain in tension εu tum** [!] b1 177.2 (3) 60’100 (3) 8.25 (6) 8.90 (4) 1.87 (4) s1 167.2 (3) 55’300 (3) 6.97 (3) 9.82 (2) 5.21 (2) s2 169.1 (3) 56’567 (3) 6.17 (6) 13.53 (6) 3.08 (6) s3 177.1 (3) 58’900 (3) 6.48 (6) 10.83 (6) 2.30 (6) * measured on 70 × 140 mm cylinders. ** calculated by inverse analysis of bending tests on 30 mm thick plates according to [9]. note: (2), (3), (4), (6) indicate the number of tests used to calculate the average value. table 3. measures mechanical properties for uhpc at 28 days central part of these beams was reinforced with an additional " 20 b500b steel rebar. beam b1 was designed to fail in bending and was tested in four-point bending with 3.0 m span. each beam was fabricated with a separate uhpc batch. prestress forces of 115 kn per wire were freed three days after casting. instantaneous and delayed shortening of the beams were monitored until the day of testing for a precise evaluation of prestress losses. for "s" elements, longitudinal strains on "sides b" were monitored while testing "sides a": strains very close to zero were detected, showing that the prestressing level on "sides b" was not influenced by cracking and failure that occurred on sides "a". tests on "side a" were stopped relatively early after the peak force, while tests on "side b" were pursued much further in the post-peak to verify the ductility of the beams. 4.2. materials the commercial product ductal fm was used as uhpc mixture for the test series, with selfcompacting consistency and 270 kg/m3 (3.4% in vol.) high-strength straight steel fibers (length lf = 14 mm, diameter df = 0.2 mm). table 3 summarizes the average 28 days mechanical properties for each batch. measured values are comparable but slightly higher than those assumed for preliminary design. however, relatively low values of the elastic limit in tension, fute, were measured for some specimen. this could be partly explained by the surface grinding procedure which was applied to 30 mm thick plates according to [9], which may have damaged these specimens. 4.3. test results all load tests were carried out at an age of uhpc comprised between 28 and 34 days. table 4 summa492 vol. 33/2022 prestressed uhpc beams beam/ batch web opening maximal measured force numerically predicted max. force maximal measured shear maximal measured moment expected strength (see tab. 2) required strength (see tab. 2) shear/moment at 1st detectable* crack b1 no 86.4 kn 43.2 knm 46.3 knm 43.5 knm 39.0 knm s1-a no 146.3 kn 158.8 kn 73.2 kn 69.7 kn 34.7 kn 54.9 kns1-b 174.3 kn 87.2 kn 69.7 kn 34.7 kn 75.2 kn s2-a "96 149.6 kn 151.0 kn 74.8 kn 34.7 kn 35.0 kns2-b 151.3 kn 75.7 kn 34.7 kn 44.9 kn s3-a 96 × 150 139.7 kn 127.8 kn 69.9 kn 34.7 kn 40.1 kns3-b 135.4 kn 67.7 kn 34.7 kn 45.2 kn * cracks detected either by visual inspection or by digital image correlation. table 4. summary of test results. figure 5. measured force-displacement curve and crack pattern at failure for beam b1. rizes the maximal applied force and the corresponding maximal shear force and bending moment for each test. these values are compared to expected and required average strength values according to preliminary design (see table 2) and to strength predictions obtained with non-linear numerical modeling (see section 5). figures 3, 5, 6 and 7 present force versus mid-span displacement curves (at two different displacements scales) and pictures of the crack patterns at failure for each beam. based on observations and measured values data, the main experimental findings are summarized: • all beams showed the expected failure mode • for all beams, including beams with web openings, the measured strength is equal or higher than the required values (see tables 1 and 2). for beams "s", the location of the critical shear crack was strongly influenced by the position where the additional fă20 rebar ends • test results of sides "a" and "b" are practically identical for beams s2 and s3, but show aăsignificant difference for beam s1 (figure 5). this difference could be explained by the accidental presence of a failure plane with particularly poor number or orientation of fibers on side "a". more detailed analyses are however required to explain this behavior • by taking as a reference the average strength of beams s1-a and s1-b, the reduction of shear strength due to presence of transversal web openings is in any case less than 20% • for all beams, ultimate behavior is extremely ductile, with the progressive formation of a collapse mechanism showing very significant postpeak residual strength • cracks were only detected (tableă4, note 1) at relatively high load levels, indicating that uhpc beams will be uncracked or micro-cracked in service. 5. non-linear finite element modeling to gain a deeper understanding on test results and to carry out further parametric analyses, the structural behavior of type "s" beams was modeled with a non-linear finite element software [14]. the software solution implements a coupled tensile fracture, compressive plasticity model and allows to define strain hardening/strain softening constitutive laws within a smeared crack modeling approach. for the sake of simplicity, a unique strain-hardening tensile law was used for all the beams, with fu te, fu tm and εu tu calculated as the average of all the bending tests carried out for batches s1, s2, s3 and b1 (figure 4, black). for the softening part, a stress-crack opening softening law according to [15] was assumed, with zero stress for a crack opening equal to lf /2. to consider 493 d. redaelli, j. moix, a. muresan et al. acta polytechnica ctu proceedings figure 6. tensile stress-strain and stress-crack opening laws for modeling. figure 7. measured vs simulated force-displacement curve and crack pattern at failure for beam s1. 494 vol. 33/2022 prestressed uhpc beams figure 8. measured vs simulated force-displacement curve and crack pattern at failure for beam s2. figure 9. measured vs simulated force-displacement curve and crack pattern at failure for beam s3. 495 d. redaelli, j. moix, a. muresan et al. acta polytechnica ctu proceedings the less favorable alignment of fibers in beams than in thin bending plates, the post-cracking part of the tensile law was scaled on the stress axis by aăfactor 0.8 (figure 4, red). young’s modulus and compressive strength were set according to values given in table 3, while other parameters of the three-dimensional mechanical model were set according to state-of-the-art knowledge and are not detailed here. numerical solutions are compared to experimental ones in figures 5 to 7: despite some numerical instabilities in the post-peak and the need to validate the sensitivity of the numerical solution to variations in the numerous mechanical parameters, the experimental response is quite well reproduced both in terms of crack pattern and forces vs mid-span displacements. 6. conclusions and future work this paper presented the design, testing and preliminary modeling of uhpc beams that were conceived as a possible technical implementation of a new modular, reusable and extremely versatile structural system for two-way slabs of residential and office buildings. experimental and modeling results demonstrate that uhpc beams can be designed to meet the strength and serviceability requirement of the new modular structural system with the imposed limitations of height and with a relatively light crosssection. thanks to the tensile contribution of the fibers, the presence of web openings of significant size only has a limited influence on the mechanical performances of the beams, thus allowing horizontal technical shafts to pass through the structural thickness of the slab. in the next steps of the research program, the economic and environmental impact of the system will be evaluated and compared to other possible implementations of the new modular system (e.g. timber, steel or mixed-construction) and to traditional, nonreusable constructive systems for two-way slabs. the sensitivity of finite element modeling with respect to variation of the input mechanical parameters will be checked, then the model will be used as a tool for parametric analysis and for further optimization of the proposed concept. acknowledgements the authors are very grateful to the whitbybird foundation for funding j. a.’s initial work on the original paper [7]; and to the open-oxford-cambridge arts and humanities research council (ahrc grant number ah/r012709/1) doctoral training partnership for funding the continued project. funding for a. m. came from the school of engineering and innovation at the open university references [1] european environmental agency. national footprint accounts, 2018. [2] danish environmental protection agency. building a circular future, 3rd edition, 2018. https://gxn.3xn.com/wp-content/uploads/sites/4/ 2018/09/building-a-circular-future_3rd-edition _compressed_v2-1.pdf. [3] l. pérez-lombard, j. ortiz, c. pout. a review on buildings energy consumption information. energy and buildings 40(3):394-8, 2008. https://doi.org/10.1016/j.enbuild.2007.03.007. [4] j. layke, e. mackres, s. liu, et al. accelerating building efficiency, washington: world resources institute, pp. 156, 2016. https://www.wri.org/resear ch/accelerating-building-efficiency. [5] fc. fivet. design of load-bearing systems for open-ended downstream reuse. iop conference series: earth and environmental science 225, 2019. https://doi.org/10.1088/1755-1315/225/1/012031. [6] c. fiver, j. brütting, nothing is lost, nothing is created, everything is reused: structural design for a circular economy, the structural engineer 98(1):74-81, 2020. https://www.researchgate.net/profile/jan-b ruetting-2/publication/338342629_nothing_is_los t_nothing_is_created_everything_is_reused_stru ctural_design_for_a_circular_economy/links/5e 4581da299bf1cdb9283f37/nothing-is-lost-nothing -is-created-everything-is-reused-structural-d esign-for-a-circular-economy.pdf. [7] a. muresan, j. brütting, j. canada j, et al. proc. of the iass symp. creativity in structural design ed c mueller and s adriaenssens, boston: mit, 2018. [8] swiss society of engineers and architects. sia 265 timber construction, 2012. [9] swiss society of engineers and architects. sia 2052 ultra-high performance concrete (uhpc) materials, design and construction, 2016. [10] j. hegger, g. bertram. tailor made concrete structures: new solutions for our society, crc press, p. 96, 2008. [11] r. zagon, s. matthys, z. kiss. shear behaviour of sfr-uhpc i-shaped beams. construction and building materials 124:258-68, 2016. https: //doi.org/10.1016/j.conbuildmat.2016.07.075. [12] l. moreillon, j. nseir, j. tomas, et al. rilem-fib-afgc int. sym. on ultra-high performance fibre-reinforced concrete, pp 549-556, 2013. https://www.afgc.asso.fr/app/uploads/2014/ 01/pp-549-556-moreillon-1.pdf. [13] redaelli d and nseir j 2020 fibre reinforced concrete. from design to structural applications, fib bulletin no. 95 pp 474-483. [14] v. červenka, l. jendele, j. červenka. atena program documentation theory, 2020. https://www.cervenka.cz/assets/files/atena-pdf /atena_theory.pdf. [15] d. redaelli. comportement et modélisation des éléments de structures en béton fibré à ultra-hautes performances avec armatures passives. école polytechnique fédérale de lausanne, switzerland, lausanne, ph.d. thesis, p. 285, 2009. 496 https://gxn.3xn.com/wp-content/uploads/sites/4/2018/09/building-a-circular-future_3rd-edition_compressed_v2-1.pdf https://doi.org/10.1016/j.enbuild.2007.03.007 https://www.wri.org/research/accelerating-building-efficiency https://doi.org/10.1088/1755-1315/225/1/012031 https://www.researchgate.net/profile/jan-bruetting-2/publication/338342629_nothing_is_lost_nothing_is_created_everything_is_reused_structural_design_for_a_circular_economy/links/5e4581da299bf1cdb9283f37/nothing-is-lost-nothing-is-created-everything-is-reused-structural-design-for-a-circular-economy.pdf https://doi.org/10.1016/j.conbuildmat.2016.07.075 https://www.afgc.asso.fr/app/uploads/2014/01/pp-549-556-moreillon-1.pdf https://www.cervenka.cz/assets/files/atena-pdf/atena_theory.pdf acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.32.0007 acta polytechnica ctu proceedings 32:7–14, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague considerations for brittleness in tailings jiří herzaa, ∗, ryan singhb a czech technical university in prague, faculty of civil engineering, depatment of geotechnics, thákurova 7, 166 29 prague, czech republic b hats consulting, 118 st georges terrace, pert, wa 6000, australia ∗ corresponding author: jiri.herza@fsv.cvut.cz abstract. brittleness is an important geotechnical parameter of soils as it describes the degree of reduction of the soil shear resistance after reaching the peak strength. [1] described soil brittleness by means of a brittleness index (ib), which is the ratio of the shear resistance loss to the peak shear strength. the ib has become a common soil parameter that is used as an indicator for the soil susceptibility to liquefaction. however, ib does not consider the rate at which the soil resistance reduces, the development of excess pore water pressure and it ignores the stress strain relationship. as a result the same ib can be calculated for a soil that collapses over a very small strain range and a soil that gradually reduces its shear resistance over extensive strain levels as long as both soils have similar peak and residual shear strengths. this paper discusses the root causes of the shear resistance loss and proposes a modification of the ib parameter to take account of the rate of the shear resistance loss, which will help to better depict the stress-strain behaviour of strain softening soils. this paper also distinguishes between structural brittleness and undrained brittleness and provides a clear definition of each to improve the understanding of tailings among practitioners. keywords: brittleness, critical state, shear resistance, tailings. 1. introduction following the failures of the tailings dams at samarco (fundão) and brumadinho (feijão) in brazil, brittle failure has become a topic of great interest within the geotechnical engineering community. brittle failure has been identified to be a contributing factor in the failures of many tailings storage facilities, as described by [2–4] and others. the post-failure review of the feijão dam [4] identified that the sudden failure of the dam was caused in part by tailings exhibiting brittle behaviour. of concern to the mining industry was the finding that there were no apparent signs of distress prior to failure, which is in fact a fundamental characteristic of brittle failure. it should be noted that the lack of indication of distress was not due to a lack of monitoring instrumentation for the dam which featured 113 piezometers, 14 surface survey markers, six inclinometers and 50 monitored seepage points. the industry’s concern regarding the presence of brittle materials within and supporting tailings dams, particularly in those featuring upstream raises is evident in the requirements of the newly published global industry standard on tailings management ("the global standard"), which includes requirement 4.6 that reads "identify and address brittle failure modes with conservative design criteria."(global tailings review, 2020). this requirement is also reflected in several other design guidelines, such as australian national committee on large dams [5] guideline on tailings dams [6], which provides recommendations for conservative design assumptions if materials are found to be susceptible to static liquefaction, which is noted to be a brittle subset of contractive materials. other forthcoming publications, guidelines and papers – including the forthcoming international commission on large dams (icold) bulletin on tailings dam safety (icold, in progress) and the international council on mining and metals (icmm) tailings management good practice guide (icmm, 2021) identify the identification of brittle failure modes as a critical evaluation factor for dam safety that should be a primary consideration in the design and review of tailings storage dams. despite the recent – and perhaps overdue – emphasis on brittleness and its importance to tailings dams, the meaning of brittleness and methods to characterize it currently lacks certain aspects of clarity and alignment with how it is not being used. further, the concept may not be fully understood by some professionals involved in evaluating and designing these facilities. generally, the use of the term "brittle" in the context of recent tailings dam failures and in the various standards and guidelines that refer to brittle behaviour alludes to a sudden loss of shear strength without signs of prior deformation. however, the currently available definition for soils as proposed by bishop (1967) does not fully align with this interpretation, namely the aspect of suddenness. without a formal definition of brittleness that reflects 7 https://doi.org/10.14311/app.2022.32.0007 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en jiří herza, ryan singh acta polytechnica ctu proceedings the intended use, the meaning or measure of terms like "brittle material or brittle failure mode" is left to the individual interpretation of the reader. 2. brittleness index bishop’s index (bishop, 1967) relates the peak shear strength (τp) to the residual shear strength (τr) as follows: ib = τp − τr τp (1) the index provides a sense of the magnitude between the peak and residual shear stresses, with a higher index indicating a greater difference between the peak and residual resistance of a material. the ib proposed by bishop does not consider the strain (deformation) levels at which a brittle failure is initiated or the strain at which the residual strength is achieved. the importance of strain has been well recognised in the field of rock mechanics, where it has been identified that the brittleness of a rock mass provides insight into how the rock mass will behave for various applications, e.g. hydraulic fracturing , drilling rate, dust generation, etc. [7]. over a dozen brittleness indices have been developed for use in rock mechanics to capture the various aspects of material strength loss, including consideration of the strain range within which the loss occurs [8]. [1] recognised that ib did not consider the suddenness of the shear strength loss and thus he also introduced the concept of brittleness in terms of the additional energy required to progress the shear failure from the peak (ep) to the residual state (ef ) as shown in figure 1. figure 1. an energy parameters for brittle soils (bishop, 1967). while the rupture index (introduced in figure 1) considered the suddenness of the shear strength loss, [1] recognised that the shape of the stains-stress line past the failure is strongly dependant on numerous parameters including test type, test procedure, sample preparation, geometry etc. as a result the rupture index has not been adopted by the geotechnical community and the strain considerations in the brittleness assessment for soil have not been addressed in the literature. in addition, bishop’s work was focused on clays, whereas tailings frequently (but not exclusively) occur as cohesionless silts and sands. while the magnitudebased brittleness index may be useful in both cohesive and cohesionless materials, the mechanism of strength loss is typically different between the two classes of materials, as discussed further below, and the strain levels at a which brittle failure occurs may vary significantly. [4] thus proposed renaming bishop’s brittleness index as the "strength loss index" because it considers solely the strength and not the strain aspects of brittleness. 3. different types of brittleness the research leading to the definition of bishop’s ib was primarily focused on the behaviour of plastic clays. the shear strength reduction depicted in bishop’s work was the result of the clay particles reorientating along the developing shear plane. [1, 2] drew attention to the decreasing shear resistance of plastic clays caused by the development of continuous bands of strongly orientated particles, which resulted in the strength of the clays lowering to a residual value. [2] observed the development of strains and displacement during simple shearing of the kaolin clay fabric. figure 2 (left) shows successive sections of kaolin clay, which were interrupted at various stages of shearing. each section has been labelled (e.g. v.2, v.3, etc.), to show the position of the section on the adjacent stress-strain curve. as can be seen, the development of structural strains and displacements of the clay structure coincides primarily with an increase in the peak strength as the original clay structure rotates (v.3). in the post-peak phase of shearing, discontinuities of the structure occur with progressive reorientation of the clay particles occurring at discontinuities (figure 2, right). once the clay minerals had reorientated along the direction of strain, the residual shear strength of the clay was reached. the originally proposed brittleness index therefore relates the loss of shear strength of the material to the structural change of the material caused by straining, which is an irreversible phenomenon. strain-induced shear strength loss also occurs in granular materials where the soil particle and intraparticle bonding can be compromised by shearing. in sandy soils, the structure of the sand and sand-like particles can be crushed and grinded as the particles move over and around each other during shearing. this change to the particles and intra-particle contacts can result in a reduction of the frictional resistance of the soil during shearing. this loss in shear strength is intensified for cemented granular materials in which straining can break the cementitious bonds resulting 8 vol. 32/2022 considerations for brittleness in tailings in a sudden (and potentially large) drop in the shear strength of the material. figure 3 shows an example of cemented aggregates of bauxite residue in which sandsite particles are composed of much smaller particles interconnected by iron oxide and aluminium hydroxide minerals. although it is acknowledged that the properties of soils change in time, the structural changes and the peak and the residual strength can be considered, for the purpose of this paper, as lasting properties of the soil induced by straining. in line with the original bishop’s brittleness concept, the shear strength in the context of this paper is considered to be a soil property that depends primarily on the soil composition and structure. in comparison, shear resistance is hereby defined as the ability of a soil to withstand shear stress under a given set of conditions. it must be emphasised that a generic function of the shear resistance has numerous parameters, some of which may still be unknown. the influencing parameters include the void ratio, soil composition, stress history, level of structuration, temperature, strain, strain rate and structure. the limitations of our testing and computing methods do not allow us to fully quantify the influence of these parameters; thus, the shearing relationships are simplified, such as in the mohr-coulomb equation which is the most common expression of the shear resistance of soils. the shear strength of a soil in the mohr-coulomb equation is expressed by the indices of an effective friction angle (φ′) and an effective cohesion (c′) and the shear resistance of the soil at failure (τf ) is expressed as: τf = c′ + σ′n tanφ ′ (2) where σ′n represents the normal effective stress acting on the shear surface, the sole variable in the relationship. to illustrate the concept of shear strength as a soil property and shear resistance as a function of stress conditions, the φ′ value has a definite minimum value that is greater than zero. however, the normal effective stress could reduce to zero due to significant increase in pore water pressures. using the example of the feijao tailings dam, the cause for the loss of shearing resistance was the rapid and significant increase of pore water pressures, which reduced the effective normal stress variable σ′n to zero or near zero and resulted in a liquefaction of the tailings. the generation of excess pore water pressure was caused by the contractive nature of the tailings in which the collapse of the loose aggregate assembly transferred the stress from the interparticle interfaces to the water. it is noted that the breakdown of the iron oxide bonds (similar to those shown in figure 3 reduced the effective cohesion value of the feijao tailings. however, it is unlikely that the frictional resistance between particles, represented by c′ and φ′, changed in any substantial way and was the root cause of the failure. because of the major differences between the described processes leading to the reduction of the soil’s ability to resist shear stress, it is proposed that shear resistance loss caused the irreversible change of the soil’s structure as described by [1] and others be referred to as structural brittleness. the sudden loss in a soil’s shear resistance caused by excess pore water pressure should then be referred to as undrained brittleness. it is recognised that both types of brittleness can occur simultaneously, as structural brittleness can also lead to undrained brittleness and many soils can lose shear resistance due to a combination of shear strength reduction and excess pore water pressure generation. however, an understanding of the different underlying phenomena leading to a shear resistance loss is crucial because the minimum recommended factors of safety that are currently used to prevent embankment failures do not account for misconceptions of the underlying geotechnical processes [5]. 4. structural brittleness and the rate of strength loss as discussed earlier, the current brittleness index alone lacks a measure of the suddenness of a material’s resistance or strength loss, with bishop’s rupture index indirectly providing a method to assess the strain over which shear resistance is lost. a more direct consideration of the strain range over which the loss of shear strength occurs was assessed by estimating the rate of the greatest reduction in shear strength after the peak shear strength is reached normalised by the peak shear strength, as shown in figure 4 with the stress stain data adopted from [2] (v7 plot in figure 2, with the data translated to si units). this calculated value, called the brittleness slope (bs), provides for the aspect of "suddenness" of strength loss and expresses the percentage of the peak shear strength loss per 1 mm of shear strain. to illustrate the application of the bs, the index was calculated for sands and clays with available direct shear test data. of particular interest was the application of the index for sands of varying cementation, density in the form of relative density/void ratio and normal effective stress. direct shear test data for sands was obtained from [10, 11]. the tests were completed on uncemented samples and with an addition of 1% and 2% of cement and at different normal effective stresses. additional ring shear tests on interparticle bonded bauxite residue (figure 3) were also completed for the purpose of this study. for a comparison to the above non-plastic soils direct shear test data for clays was also included, based on direct shear tests for a clay material presented by [12] and the shear box test results from [2]. bishop’s brittleness index and the brittleness slope for the referenced tests are shown in figure 5. 9 jiří herza, ryan singh acta polytechnica ctu proceedings as can be seen, the development of structural strains and displacements of the clay structure coincides primarily with an increase in the peak strength as the original clay structure rotates (v.3). in the post-peak phase of shearing, discontinuities of the structure occur with progressive reorientation of the clay particles occurring at discontinuities (fig. 2, right). once the clay minerals had reorientated along the direction of strain, the residual shear strength of the clay was reached. the originally proposed brittleness index therefore relates the loss of shear strength of the material to the structural change of the material caused by straining, which is an irreversible phenomenon. strain-induced shear strength loss also occurs in granular materials where the soil particle and intra-particle bonding can be compromised by shearing. in sandy soils, the structure of the sand and sand-like particles can be crushed and grinded as the particles move over and around each other during shearing. this change to the particles and intra-particle contacts can result in a reduction of the frictional resistance of the soil during shearing. this loss in shear strength is intensified for cemented granular materials in which straining can break the cementitious bonds resulting in a sudden (and potentially large) drop in the shear strength of the material. fig. 3 shows an example of cemented aggregates of bauxite residue in which sand-site particles are composed of much smaller particles interconnected by iron oxide and aluminium hydroxide minerals. although it is acknowledged that the properties of soils change in time, the structural changes and the peak and the residual strength can be considered, for the purpose of this paper, as lasting properties of the soil induced by straining. in line with the original bishop’s brittleness concept, the shear strength in the context of this paper is considered to be a soil property that depends primarily on the soil composition and structure. in comparison, shear resistance is hereby defined as the ability of a soil to withstand shear stress under a given set of conditions. fig. 2: sequence of kaolin clay structures sheared normal to original clay fabric (left) and sequential development of discontinuity with arrow indicates direction of shearing (right) (morgenstern & tchalenko, 1967) figure 2. sequence of kaolin clay structures sheared normal to original clay fabric (left) and sequential development of discontinuity with arrow indicates direction of shearing (right) [2]. fig. 3: cemented aggregates of bauxite residue (herza, 2020) it must be emphasised that a generic function of the shear resistance has numerous parameters, some of which may still be unknown. mitchell and soga (2005) listed several influencing parameters including the void ratio, soil composition, stress history, level of structuration, temperature, strain, strain rate and structure. the limitations of our testing and computing methods do not allow us to fully quantify the influence of these parameters; thus, the shearing relationships are simplified, such as in the mohr-coulomb equation which is the most common expression of the shear resistance of soils. the shear strength of a soil in the mohr-coulomb equation is expressed by the indices of an effective friction angle (φ’) and an effective cohesion (c’) and the shear resistance of the soil at failure (tf) is expressed as: where σ’n represents the normal effective stress acting on the shear surface, the sole variable in the relationship. to illustrate the concept of shear strength as a soil property and shear resistance as a function of stress conditions, the φ’ value has a definite minimum value that is greater than zero. however, the normal effective stress could reduce to zero due to significant increase in pore water pressures. using the example of the feijao tailings dam, the cause for the loss of shearing resistance was the rapid and significant increase of pore water pressures, which reduced the effective normal stress variable σ’n to zero or near zero and resulted in a liquefaction of the tailings. the generation of excess pore water pressure was caused by the contractive nature of the tailings in which the collapse of the loose aggregate assembly transferred the stress from the interparticle interfaces to the water. it is noted that the breakdown of the iron oxide bonds (similar to those shown in chyba! nenalezen zdroj odkazů.) reduced the effective cohesion value of the feijao tailings. however, it is unlikely that the frictional resistance between particles, represented by c’ and φ’, changed in any substantial way and was the root cause of the failure. because of the major differences between the described processes leading to the reduction of the soil’s ability to resist shear stress, it is proposed that shear resistance loss caused the irreversible change of the soil’s structure as described by bishop (1967) and others be referred to as structural brittleness. the sudden loss in a soil’s shear resistance caused by excess pore water pressure should then be referred to as undrained brittleness. it is recognised that both types of brittleness can occur simultaneously, as structural brittleness can also lead to undrained brittleness and many soils can lose shear resistance due to a combination of shear strength reduction and excess pore water pressure generation. however, an understanding of the different underlying phenomena leading to a shear resistance loss is crucial because the minimum recommended factors of safety that are currently used to prevent embankment failures do not account for misconceptions of the underlying geotechnical processes (herza et al 2017). figure 3. cemented aggregates of bauxite residue [9]. figure 4. brittleness index for strength loss. 10 vol. 32/2022 considerations for brittleness in tailings a "zone of high brittleness" has been overlaid on the structural brittleness plane to identify areas where ib is greater than 0.4. this ib threshold has been adopted from [4]. it should be noted that this value of 0.4 was selected as a threshold by robertson as available data for flow liquefaction failures indicated a correlation of brittleness and liquefied undrained shear strength ratios. in reviewing the conclusions in [4], it is likely that structural brittleness and undrained brittleness has been combined for several reasons, including the testing being carried out in undrained conditions leading to the generation of excess pore water pressures. nonetheless, where ib values greater than 0.4 are observed it indicates that particular attention should be given to the potential for some form of brittle behaviour. figure 5. structural brittleness plane. the analysed test results indicated that: • brittleness of clean uncemented sand is smaller compared with the brittleness of the cemented materials and it increases with the increasing sand density as expected. • increasing cement content increases the peak shear strength, the dilatancy and the ib and bs values. it also results in the peak shear strength being achieved at lower shear displacements. • increasing normal stresses by 100% supresses the dilatancy and the ib and bs values of the samples with 2% cement content. however, the increase of normal stress had only a very little impact on the samples with 1% of cement added. • the spread of the points in figure 5 is most likely a result of the non-linear relationships between the sand density, cement content and the applied effective stress. in addition, several of the presented stress paths did not converge to the same residual shear strength at the same normal stress, indicating that the shearing may not have destroyed all cemented bonds within the sand samples. • the test results of the bauxite residue were showing brittleness caused by gradual destruction of the intra-particle bonds only on comparatively denser samples where the loss of shear strength was accompanied by the sample contraction. this behaviour contrasted with the behaviour of the cemented and uncemented compacted sands that exhibited dilation after reaching the peak shear strength. • the bauxite residue samples reached the peak shear strength values at higher displacement compared to the artificially cemented sands. for clays, it was noted that the range of brittleness varied significantly and the brittleness slope reduces with an increasing brittleness index. however, it is likely that due to the low permeability of the tested clays some of the measured loss of shear resistance was due to the development of shear-induced porewater pressures. in relation to the brittle behaviour identified in recent tailings dam failures, structural brittleness of the tailings was not the primary trigger mechanism, nor can it necessarily indicate that undrained brittleness is likely. as expected, the test results of loose uncemented clean sands did not show structural brittleness as the samples contract during drained shearing and the tighter packing of sand particles made shearing harder. conversely the same sands, but in a denser state, indicated moderate structural brittleness as they dilated under shearing. if saturated and tested in undrained (constant volume) conditions, the dilative sands would gain shear resistance due to reduced pore pressures generated during shearing. in contrast the loose sands tested in undrained conditions would show a loss of shear resistance during shearing due to shear induced excess pore water pressures. it can therefore be seen that for uncemented clean sands, the tendency or likelihood for structural and undrained brittleness are opposed. in cemented sands however, the structural brittleness may indicate the potential for undrained brittleness and liquefaction. in the case of drained conditions, structural brittleness can be observed for loose and cemented sand, as the cementitious bonds break, resulting in a collapse of the soil, or particle structure followed by further contraction. this behaviour was observed in the test results of the bauxite residue sample. under undrained conditions, this tendency for contractive behaviour would not be able to develop and would instead translate into generation of excess pore water pressures, a reduction in normal effective stress and loss of shear resistance potentially to the point of liquefaction. as such, structural brittleness can be used as a screening assessment for the potential of undrained brittleness and even liquefaction for loose and cemented sands. this relationship could be made even clearer, with a three-dimensional path of shear stress, shear strain and volumetric change (reduction). the volumetric change is perhaps a better indication of structural change than a void ratio for cemented materials as the change of the structure (breakage of particles) may not necessarily be reflected by the change of the void ratio. 11 jiří herza, ryan singh acta polytechnica ctu proceedings 5. undrained brittleness as discussed earlier, the undrained brittleness refers to the loss of shear resistance of a soil due to the straininduced pore water pressures. this type of brittleness is caused by a rapid increase of pore water pressures when the soil structure decays in so-called undrained conditions. such conditions occur in contractive, fully or partially saturated soils when the strain-induced porewater pressures do not concurrently dissipate. under such conditions, the loss of shear resistance is proportional to the reduction in effective confining stress in the soil. with the available testing apparatus, given excess porewater pressure generated in the shear zone cannot be directly measured using the current laboratory equipment, it is being inferred from observations made at the boundaries of the sample. together with the other influencing factors discussed earlier, this makes the conventional strain-stress diagrams unsuitable for description of the undrained brittleness. instead, the concept of critical state soil mechanics (cssm) provides an indication of the likelihood of a brittle failure, and this indeed has become common practice in the tailings industry. the use of the state parameter to assess a soil’s likelihood of brittle failure is based on the distance between the current state of the soil from its critical state, as expressed by a critical state line (csl). the csl is a locus of critical states, at which continuous shearing occurs without a change of mean effective stress (p′), deviatoric stress (q) and void ratio (e) [3]. due to its dependency on the void ratio as a measure of the soil structure, the cssm concept is not directly applicable to all materials, in which the orientation of aggregates and/or cementation plays an important role such is the case of the kaolinite clays and cemented sands discussed earlier in this paper. the csl of a soil is commonly depicted as a relationship between mean effective stress and void ratio, with any point on the plane representing a soil’s state prior to shearing. an example of a csl is shown in figure 6 [3]. per cssm, any soil being sheared will tend towards the csl as the soil structure decays from the initial state to the critical state. during undrained shearing of a soil, volumetric changes are prohibited and the shear stress path towards the csl is horizontal with respect to the csl plot as shown in figure 6. figure 6. csl [13]. the pre-sheared state of the soil is often described using a state parameter, ψo, which is the difference in the pre-sheared state and the critical state void ratios at the same mean effective stress (figure 6). a stress path following the vertical direction of the state parameter ψo can be obtained from a direct shear tests at a constant normal stress undertaken in perfectly drained conditions. however, as the undrained brittleness is caused by reduction of the mean effective stress in undrained shearing, the state parameter ψo alone is not sufficient to capture the potential for brittle failure. the susceptibility to brittle failure, rather, is proportional to the horizontal distance between the pre-sheared mean effective stress (p′o) and the effective stress at the csl at constant p′cs) which is equivalent to the ratio of the state parameter ψo and the csl slope (ψo/λe) as shown in figure 6. a greater ψo/λe indicates a greater potential for generation of excess pore water pressures during undrained shearing and thus a greater likelihood of brittle failure. [13] presented results of critical state lines for various tailings, which demonstrated that the csl cannot be reliably constructed from basic soil characterisation tests and csl specific testing is crucial to understand the potential for undrained brittleness. 6. discussion and conclusions this paper proposes to clarify terms for brittleness to align them with the use of the term brittle as it pertains to a soil exhibiting a sudden loss of shear strength without signs of prior deformation, particularly in the context of recent tailings dam failures. in this context the two following forms of brittle behaviour are identified: • structural brittleness – a loss of a soil’s shear strength caused by irreversible change to a soil’s structure during shearing, that can occur in clays and cemented soils • undrained brittleness – a loss in a soil’s shear resistance due to a change in the stress field applied to a material, often due to generation of excess pore water pressures the structural brittleness can be captured by bishop’s ib using conventional strain-stress plots although the rate of the shear strength loss can be 12 vol. 32/2022 considerations for brittleness in tailings treated with caution due to its dependence on the testing procedure. modern investigation techniques such as sem can be successfully applied to reveal the nature of the soil structure at the microscopic level and thus its susceptibility to structural brittleness. the undrained brittleness is the result of excess pore water pressure generated during undrained shearing and the use of ib estimated from stress-strain plots is considered insufficient. although the ib can used as a screening tool, the undrained brittleness should be considered using the cssm concept. a state parameter ψo greater than -0.05 (a dashed parallel line to the cs, in figure 7 is commonly taken as an indication of contractive soils that may be susceptible to brittle failure. rather than using ψo alone, the likelihood of brittle failure should be expressed by the ψo/λe ratio with increasing values indicating a higher likelihood of undrained brittleness. the potential susceptibility to undrained brittleness is indicated by shading of zones in figure 7. because not all contractive soils are susceptible to brittle failure, further delineation of the zones may be appropriate using the ψo/λe value. as figure 7 shows a log-linear relationship between the mean effective stress and the void ratio, the potential boundary of undrained brittleness may by a curved line as indicated in figure 7. the authors emphasise that the potential boundary of undrained brittleness is indicative only and it is not based on any testing. the structural brittleness can be captured by bishop’s ib using conventional strain-stress plots although the rate of the shear strength loss can be treated with caution due to its dependence on the testing procedure. modern investigation techniques such as sem can be successfully applied to reveal the nature of the soil structure at the microscopic level and thus its susceptibility to structural brittleness. the undrained brittleness is the result of excess pore water pressure generated during undrained shearing and the use of ib estimated from stress-strain plots is considered insufficient. although the ib can used as a screening tool, the undrained brittleness should be considered using the cssm concept. a state parameter ψo greater than -0.05 (a dashed parallel line to the cs, in fig. 7is commonly taken as an indication of contractive soils that may be susceptible to brittle failure. rather than using ψo alone, the likelihood of brittle failure should be expressed by the ψo/λe ratio with increasing values indicating a higher likelihood of undrained brittleness. the potential susceptibility to undrained brittleness is indicated by shading of zones in fig. 7. because not all contractive soils are susceptible to brittle failure, further delineation of the zones may be appropriate using the ψo/λe value. as fig. 7 shows a log-linear relationship between the mean effective stress and the void ratio, the potential boundary of undrained brittleness may by a curved line as indicated in fig. 7. the authors emphasise that the potential boundary of undrained brittleness is indicative only and it is not based on any testing. fig. 7: csl and critical zones beyond considering this state parameter delineation, further definition can be added to a csl plot in order to better define the area of concern. the authors propose that the assessment should focus on the range of mean effective stresses that the soil can be exposed to under the current and future loading conditions. this range includes the current conditions at relevant depths and the pressures imposed by potential development and closure of the tsf. this is important because  soils that would only exhibit undrained brittleness under mean effective stresses well outside the operation stress range can be defensibly deemed not to pose a concern for the structure.  soils that are dilative in the current state can become contractive and potentially susceptible to undrained brittleness if the mean effective stress increases due to the future tsf development or closure. operating stress range in structural zone of tsf ? ? ? ? ? ? ? ? ? potential boundary of undrained brittleness? csl ψ0 = 0.05 figure 7. csl and critical zones. beyond considering this state parameter delineation, further definition can be added to a csl plot in order to better define the area of concern. the authors propose that the assessment should focus on the range of mean effective stresses that the soil can be exposed to under the current and future loading conditions. this range includes the current conditions at relevant depths and the pressures imposed by potential development and closure of the tsf. this is important because: • soils that would only exhibit undrained brittleness under mean effective stresses well outside the operation stress range can be defensibly deemed not to pose a concern for the structure. • soils that are dilative in the current state can become contractive and potentially susceptible to undrained brittleness if the mean effective stress increases due to the future tsf development or closure. the authors propose that consideration of appropriate mean effective stresses be included in the assessment of materials susceptible to undrained brittleness in the design and assessment of all civil structures, including tsfs. to support this, the derivation and application of tailings-specific csls are critical, as recommended by [13]. it should also be noted that the use of the state parameter to identify a material’s susceptibility to brittle failure is dependent on a material’s void ratio, and any materials that do not exhibit a meaningful change in void ratio during brittle failure, such as cemented sands or structured clays, cannot be appropriately assessed within this framework. thus, further research into other appropriate measures of a material’s structure may allow a wider assessment of brittle behaviour in soils and tailings, in particular. future work extending from this paper include further definition and characterisation of the causes and mechanisms of undrained brittleness, as well as some proposed index to measure the susceptibility of a material to feature undrained brittleness and thus a formal potential boundary of undrained brittleness. this will allow a characterisation of materials that may be susceptible to undrained brittleness, which aligns with the principles of the performance based, risk informed approach to storage facility safety. acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs21/045/ohk1/1t/11. references [1] a. bishop. progressive failure with special reference to the mechanism causing it. 1967. proceeding of the geotechnical conference, oslo, sweden. vol 2., page 142, 150. [2] n. morgenster, j. tchalenko. microscopic structures in kaolin subjected to direct shear. 1967. j.s. geotechnique 17: 309-328. [3] m. jefferies, k. been. soil liquefaction – a critical state approach. 2016. taylor & francis, london. [4] p. robertson. evaluation of flow liquefaction: influence of high stresses. 2017. proceedings from 3rd international conference on performance based design in earthquake geotechnical engineering. [5] j. herza, m. ashley, j. thorp. "factor of safety? – do we use it correctly?". 2017. ancold 2017 conference, october 26-27, 2017. hobart, tas, australia. [6] guidelines on tailings dams – planning, design, construction, operation and closure – revision 1. 2019. ancold. july 2019. [7] h. zhou, j. lu, y. jiang, f. meng. a new rock brittleness evaluation index based on the internal friction angle and class i stress-strain curve. 2018. rock mechanics and rock engineering vol. 51. 13 jiří herza, ryan singh acta polytechnica ctu proceedings [8] j. jiang, w. dong, h. xinping, d. shuai. relationship between a brittleness index and crack initiation stress ratio for different rock types. 2020. advances in civil engineering, vol. 2020, article id 8091895, 12 pages. [9] j. herza, d. jirasko, i. vanicek. the influence of aggregates on tailings behaviour. 2020. proceedings of the wmess conference, prague, september 2020. [10] y. amini, a. hamidi, e. asghari. shear strength characteristics of an artificially cemented sand-gravel mixture. 2013. challenges and recent advances in geotechnical and seismic research and practices, chengdu, china. october 25-27, 2013. [11] a. simoni, g. houlsby. the direct shear strength and dilatancy of sand-gravel mixtures. 2006. geotechnical and geological engineering. 24. 523-549., https://doi.org/10.1007/s10706-004-5832-6. [12] m. heidemann, l. bressani, j. flores. residual shear strength of a residual soil of granulite. 2020. soils and rocks, são paulo, 43(1): 31-41. [13] k. smith, r. fanni, p. chapman, d. reid. critical stae testeing of tailings – comparison between various tailings and implications for design. 2019. proceeding of tailings and mine waste 2019. november 17-20, vancouver, canada. 14 https://doi.org/10.1007/s10706-004-5832-6 acta polytechnica ctu proceedings 32:7–14, 2022 1 introduction 2 brittleness index 3 different types of brittleness 4 structural brittleness and the rate of strength loss 5 undrained brittleness 6 discussion and conclusions acknowledgements references acta polytechnica ctu proceedings doi:10.14311/app.2020.28.0023 acta polytechnica ctu proceedings 28(0):23–31, 2020 © czech technical university in prague, 2020 available online at http://ojs.cvut.cz/ojs/index.php/app methods of xs data preparation for geometry with fuel dummy pavel suk czech technical university in prague, faculty of nuclear science and physical engineering, department of nuclear reactors, v holešovičkách 2, 180 00 prague 8, czech republic correspondence: sukpave2@fjfi.cvut.cz abstract. 3d deterministic core calculation represents important category of the nuclear fuel cycle and safe nuclear power plant operation. the appropriate solution was not published yet. data preparation process for non-fuel elements of the core represents the challenge for scientists. this report briefly introduce the problem of the data preparation process and gives the information about new input format for macrocode parcs (pmaxs). the best homogenization process approach is to prepare data in infinite lattice cell for fuel assemblies, which are placed next to the another fuel assembly. data for fuel assembly located next to the non-fuel region are better with preparation in the real geometry with the real boundary conditions. results of the neutron spectra study show that the pmaxs file format is well prepared for the 2 group calculation, but it is not well prepared for the multigroup calculations, however the xsec file format still gave reasonable results. keywords: eigenvalue, flux, fuel dummy, homogenization, scale, serpent, parcs. 1. data preparation process and reflector calculation nuclear power plant (npp) cores are in general too large to calculate them with exact approach of the transport equation in acceptable calculation time. based on this fact, the data preparation process and full core calculation via deterministic macrocodes are still remaining comprehensive parts of the fuel loading patterns preparation. these two processes are, in general, connected together, but in the reality, the complexity of this problem is replaced by the easier and faster approach. this approach is based on the separation of the data preparation process for each part of the core and 3d full core deterministic calculation. the research reactor cores are sometimes also calculated with the simplified methods. especially the fuel loading pattern in the high power research reactors can be designed via simplified homogenization approach. the total neutron-termohydraulic solvers are also based on this methodology. therefore the homogenization process is also important during the analysis of research reactors. the data preparation process for the fuel elements is historically based on the exact transport simulation of the fuel assembly with reflective boundary conditions [1]. the main complication of this approach is fact that the fuel assemblies are not in the infinite medium environment during the 3d calculation. the reactor core does not consist of the same fuel assemblies, some flux tilt inside the fuel assemblies can be presented. some advantages of this approach can be found in the rehomogenization process, which is described in papers [2] or [3]. the research reactors and npps are almost always operated in the critical state, but during the data preparation process, the infinite medium of the the fuel assemblies can be sub-critical or super-critical in case of their enrichment and geometry. some solution of these issues can be found in the b1 approximation, described bellow or in the article [1]. the comparison of b1, p1 and casmo methods can be found in the paper [1]. the data preparation process for non-fuel elements represents more challenge problem. the optimal solution of that problem has not been found yet, however some approaches are developed for data preparation. the main complication of the data preparation process for the non-fuel elements is based on the non-leakage medium and impossibility to calculate exact neutron spectrum during the data preparation process. the most codes are based on the diffusion solver, but the diffusion theory is not met near the boundaries [4]. data for non-fuel regions, or reflector regions can be prepared by various approaches. for instance 1d approach or more dimension approach [5]. for better criticality and power distribution prediction, the macroscopic data are sometimes optimized. the main complication of this approach is in the high calculation time. next complication of the fuel loading pattern optimization is in the uniqueness of the pattern. the optimization should to be made for each pattern separately. the optimization method can be found, for instance, in the paper [6]. numerous details of data preparation for nodal calculation codes can be found also in the paper [7]. the author explain complications of the b1 method, diffusion coefficient calculation, and spatial homogenization for data preparation process there. 23 http://dx.doi.org/10.14311/app.2020.28.0023 http://ojs.cvut.cz/ojs/index.php/app pavel suk acta polytechnica ctu proceedings 1.1. macroscopic cross section preparation macroscopic cross sections (xs) are, in general, quantities which provide information about the influence of that part of the core to the calculation. xs are defined in simple approach: σi,g,r = ni · σi,g,r, (1) where ni is atomic density of the region i and σi,g,r is microscopic cross section of region i, energy group g, and reaction r. the more rigorous approach is based on the type of the calculation code. the lattice codes can be divided into the discrete energy codes and continuous energy codes. in these two cathegories, the macroscopic cross sections are generated: σi,g,r = ∑eg,max h=eg,min ni · σi,h,r · φi,h∑eg,max h=eg,min φi,h , (2) for lattice codes which calculated with the discrete energies. φi,h is the neutron flux inside region i and energy group h. for the continuous energy calculation codes, the macroscopic cross sections are calculated via: σi,g,r = ∫ eg,max eg,min ni · σi,r(e) · φi(e)de∫ eg,max eg,min φi(e)de , (3) where σi,r(e) is the microscopic cross section of the region i, reaction r, and energy e and φi(e) is the neutron flux inside region i, energy e. 1.2. b1 correction b1 correction, also called critical spectra correction, is a method used for treating the leakage of neutrons in the fuel assemblies places in the real core pattern. as was mentioned above, the 3d deterministic calculations are based on the two level approach. the macroscopic data for fuel assemblies are prepared for the fuel assemblies in the infinite lattice of the same assemblies, thus, the neutron leakage from the fuel assembly is zero. the b1 method prepares data based on the calculation of critical spectra in the fuel assembly and due to it treats the non zero neutron leakage. the main idea of the b1 approximation is to find appropriate buckling factor and thus appropriate neutron spectra for xs preparation. the b1 corrected homogenised xs are then prepared with the new neutron spectra. the buckling factor adds or removes additional neutron leakage in the fuel assemblies based on the equation: σt,gφg ± ibjg = χg ∑ h σf,hφh + ∑ h σ0s,h→gφh, (4) where σt,g is the total macroscopic cross section in energy group g, φg is the neutron flux in energy group g, b is the buckling factor, jg is the neutron current in energy group g, χg is the neutron spectra in the energy group g, σf,h is the fission macroscopic cross section in energy group h and σ0s,h→g is the zero moment of scattering macroscopic cross section in energy group g. the critical spectra is to find iteratively by increasing buckling factor. the first step is based on the calculation with b2 = 0, the second step is based on the calculation with b2 = 10−6 and the next steps are calculated with extrapolated buckling factor from the previous calculation. the whole method is well described in the paper [1] and [7]. when the fuel assemblies in the infinite lattice have the eigenvalue less than 1, the neutron leakage is decreasing, when the eigenvalue is larger than 1, the leakage is increasing, in practical sense. 2. calculation codes calculation codes are essential for nuclear reactor operation. the codes can be divided into two main groups, macrocodes and microcodes (lattice codes). macrocodes are used for the full core calculations with diffusion or simplified transport solution [8]. microcodes are used for exact full core transport calculation or data preparation for macrocodes. microcodes can be than divided into the deterministic codes [9] and stochastic codes [10]. parcs parcs calculation code is developed at purdue university [11] for 3d npp core calculations. parcs is purdue advanced neutron core simulator code which offers various calculation methods based on the diffusion solver or simplified transport solution (sp3). some of the solvers are based on the finite difference methods, others are based on the nodal methods. [8] in the study presented below, the finite difference method based on the diffusion solver (fdm, nemmg) or simplified transport solver (sp3) are analysed. all cases were calculated with parcs v3.3.1 code version. scale scale is a comprehensive calculation package which contain deterministic codes (newt, triton) and a stochastic code (keno) [9]. the deterministic codes newt and triton are commonly used for the data preparation. both these codes are based on the solution of the transport equation in multigroup approach. xs can be prepared with actual flux spectrum or with b1 approximation. triton code is able to carried out burnup calculations in contrast with newt. all cases were calculated with scale 6.2.3 code version [12]. 3. test case the model consisted of the simplified fuel assemblies and simplified fuel dummy filled with water was developed as a test case. this model is not usual, however 24 vol. 28/2020 xs data it includes region for that the data preparation is complicated. the visualisation of the test case is shown in figure 1. this geometry can be found in the research reactors, where the water can be replaced by the research equipment like irradiation channels, detectors or other devices. figure 1. visualisation of the test case model test case consists of the 3 elements types. the whole structure is placed in the infinite lattice of the same structures. fuel assembly 1 (fa1) seems to be like a fuel assembly in the infinite lattice of the same assemblies, because each boundary is connected to the fuel assembly. fuel assembly 2 (fa2) has three sides connected with the fuel assembly and one side is connected to the water fuel dummy. water fuel dummy (dum) is simulated with the same temperature and density conditions as the moderator inside the fuel assemblies (ρ = 0.71667 g/cm3, t = 578 k). fuel assemblies consist of 17×17 fuel pins. description of the material and geometry of the test case is given in table 1. two different enrichments of the fuel pins were simulated: 0.7% sub-critical system and 4.3% super-critical system. test case parameters fuel diameter 0.82 cm cladding diameter 0.95 cm pin pitch 1.26 cm fuel density 10.219 g/cm3 moderator density 0.71667 g/cm3 fuel temperature 1100 k cladding temperature 600 k moderator temperature 578 k table 1. description of the material and geometry of the test case 4. calculation many different approaches were realised during the analysis of behaviour in the test case. in the first stage, the influence of b1 approximation and noncritical flux were analysed. the next part of the study was focused on the downscatter treatment. last part of the study deals with the multigroup approach and spectra comparison. some of the diffusion codes are designed to use only downscattering, therefore the scattering cross sections are modified via: σ̂g,h = σg,h − φh φg σh,g, (5) where the σg,h, respective σh,g is scattering cross section from group g to group h, respective from h to g. φh, respective φg is the neutron flux of group h, respective g. the compliance of the simplified solution calculated via parcs was rated by the eigenvalue compliance, see equation (6) and the neutron flux values inside homogenised areas, see equation (7). the reference solution was obtained by the whole geometry simulation in the scale 6.2.3 calculation code. ∆keff = (keff,scale − keff,parcs) · 105 (6) ∆φg = √√√√ 1 n n∑ i=0 ( φig,scale − φ i g,parcs φig,scale 100 )2 (7) 4.1. two group approach the two group xs were prepared with the scale triton calculation code in this section. the data for macrocode parcs can be prepared by 2 ways, with pmaxs files and xsec files. the data from scale triton calculation code was prepared by the genpmaxs [13] program, which prepares pmaxs files for each region. the obtained results with pmaxs files showed good agreement, therefore the xsec file format was not used during two group analyses. many calculation modes and homogenization approaches were carried during this analysis: • case1 data for each different part of model was prepared by simulation in the real geometry with the real boundary conditions with b1 correction. • case1_cor case1 with downscatter correction via equation (5). • case2 data for the fuel assemblies was prepared by simulation with the reflective boundary condition, data for dum was prepared by simulation in the real geometry with the real boundary conditions with b1 correction. • case2_cor case2 with downscatter correction via equation (5). • case3 data for fa1 was prepared by simulation with the reflective boundary condition, other data was prepared by simulation in the real geometry with the real boundary conditions with b1 correction. 25 pavel suk acta polytechnica ctu proceedings • case3_cor case3 with downscatter correction via equation (5). • case4 case1 without b1 correction. • case5 case2 without b1 correction. • case6 case3 without b1 correction. • case7 data for fuel assemblies was prepared by the simulation with the reflective boundary condition with b1, data for dum was prepared by the simulation in the real geometry with the real boundary conditions without b1 correction. • case8 data for fa1 was prepared by the simulation with the reflective boundary condition with b1, data for other regions were prepared by the simulation in the real geometry with the real boundary conditions without b1 correction. 4.1.1. 0.7 % enrichment fuel enriched only with 0.7% was simulated to investigate agreement for the sub-critical core regions. the results of eigenvalue calculation and relative flux differences are shown in table 2, visualised in figures 2 and 3. modes ∆keff (pcm) ∆φ1 (%) ∆φ2 (%) case1 51.8 4.52 1.85 case1_cor 71.7 4.56 1.92 case2 91.4 3.47 1.40 case2_cor -46.0 8.60 1.92 case3 8.6 4.23 1.67 case3_cor 26.0 4.26 1.73 case4_cor 189.5 7.92 2.61 case5_cor 183.7 6.15 2.00 case6_cor 127.6 7.34 2.30 case7_cor 66.3 4.11 1.15 case8_cor 59.1 6.15 2.00 table 2. two group results of the the test case with fuel enrichment 0.7% (reference keff = 0.82853) -50 -25 0 25 50 75 100 125 150 175 200 case1 case1_cor case2 case2_cor case3 case3_cor case4_cor case5_cor case6_cor case7_cor case8_cor k e ff d if fe re n ce ( p cm ) keff relative difference in each case for 0.7 % enrichment fuel type figure 2. relative eigenvalue differences in the calculation modes with 0.7% fuel enrichment test case in case of eigenvalue, the case1, case2_cor, case3 and case3_cor were in the best agreement with the lattice code. on the opposite, the cases without b1 correction (case4, case5 and case6) were 0 1 2 3 4 5 6 7 8 9 case1 case1_cor case2 case2_cor case3 case3_cor case4_cor case5_cor case6_cor case7_cor case8_cor p d if fe re n ce ( % ) relative power difference in each case for 0.7 % enrichment fuel type g1 g2 figure 3. relative power differences in the calculation modes with 0.7% fuel enrichment test case 1 4 -2.19 0.49 4 5 -1.88 0.14 7 6 -2.19 0.49 2 3 -1.88 0.14 5 10 -11.30 -4.89 8 7 -1.88 0.14 3 2 -2.19 0.49 6 1 -1.84 0.15 9 8 -2.19 0.49 1 2 3 1 2 3 legend n type group 1 group 2 figure 4. 2d flux differences for each part of the test case in case3 and enrichment 0.7% in the worse agreement. the neutron flux comparison showed that the thermal neutron group deviates around 2% in all cases, but the fast neutron group flux difference reached up to 9% in case2_cor, case4_cor and case6_cor. the best results in case of eigenvalue and φg are in case3. 2d flux differences of case3 are shown in figure 4. figure 4 shows the best agreement in the thermal group (only 0.49%) for fuel assemblies and only 2% deviation in the fast energy group. the neutron fluxes in the fuel dummy region inside the test case were determined with worse agreement. up to 5% for thermal energy group and up to 11.5% for the fast energy group. 4.1.2. 4.3 % enrichment the nuclear fuel in the npp cores is in general enriched up to 5%. the fuel pellets with enrichment 4.3% are usually in the active length of the fuel pins and, moreover, they make the simulated system supercritical. the results of the eigenvalue and the relative flux differences are shown in the table 3, visualised in figures 5 and 6. the best agreement in case of eigenvalue were reached in the cases without b1 correction (case4_cor, case5_cor, and case6_cor). on the other hand, the fast neutron flux differences 26 vol. 28/2020 xs data modes ∆keff (pcm) ∆φ1 (%) ∆φ2 (%) case1 688.6 5.17 4.89 case1_cor 690.6 5.14 4.97 case2 767.6 5.43 4.74 case2_cor 730.6 5.49 4.87 case3 625.6 5.16 4.81 case3_cor 626.6 5.14 4.89 case4_cor 173.6 7.82 4.61 case5_cor 121.6 9.61 4.52 case6_cor 63.6 8.23 4.41 case7_cor 792.6 5.15 5.84 case8_cor 418.6 5.93 6.36 table 3. two group results of the test case with fuel enrichment 4.3% (reference keff = 1.32588) 0 100 200 300 400 500 600 700 800 case1 case1_cor case2 case2_cor case3 case3_cor case4_cor case5_cor case6_cor case7_cor case8_cor k e ff d if fe re n ce ( p cm ) keff relative difference in each case for 4.3 % enrichment fuel type figure 5. relative eigenvalue differences in the calculation modes with 4.3% enrichment fuel test case 0 1 2 3 4 5 6 7 8 9 10 11 case1 case1_cor case2 case2_cor case3 case3_cor case4_cor case5_cor case6_cor case7_cor case8_cor p d if fe re n ce ( % ) relative power difference in each case for 4.3 % enrichment fuel type g1 g2 figure 6. relative power differences in the calculation modes with 4.3% enrichment fuel test case were the highest in that three cases. the 2d flux differences of case6_cor is shown in figure 7. case6_cor model showed the best result in case of eigenvalue, with difference only 63 pcm. figure 7 shows relatively good agreement in thermal neutron flux in fuel assemblies and fast neutron flux in dum, only up to 3%. fast neutron flux in fuel assemblies and thermal neutron flux in the dum were determined with worse agreement, up to 11%. 1 4 8.63 -2.28 4 5 8.76 -3.21 7 6 8.63 -2.28 2 3 8.76 -3.21 5 10 -2.08 -10.62 8 7 8.76 -3.21 3 2 8.63 -2.28 6 1 8.78 -3.23 9 8 8.63 -2.28 1 2 3 1 2 3 legend n type group 1 group 2 figure 7. 2d flux differences for each part of the test case in case6_cor and enrichment 4.3% 4.1.3. discussion the above results showed that the data preparation based on the actual boundary condition cannot bring good results, because the parcs calculation code does not know the actual neutron leakage from the system and due to it there is lack of the informations to properly simulate the system. the best agreement in the lower enriched case was obtained with the b1 correction. this situation corresponds with the physical solution. if the solved problem is sub-critical, the b1 approximation simulates less neutron leakage to reach criticality during the data preparation process. because the each part of the test case is "locally" sub-critical, it means that the neutron spectra is changed in each part with the same direction and due to that the results are better with b1 approximation. the situation in the higher enriched case is different. the whole test case is super-critical, but the water fuel dummy is "locally" sub-critical. the b1 correction increases neutron leakage from the whole system, but the water fuel dummy is "locally" sub-critical. the b1 correction changes the neutron spectra in both regions with the same direction and due to that the neutron spectra in the water fuel dummy is very different during the xs preparation. the better results were obtained without b1 correction and it also corresponds with the above theory. 4.2. multigroup approach to better understand the behaviour of the water-fuel neutron transport in the macrocodes, there was pressure to analyse and compare multigroup neutron flux in the parcs and scale code systems. the fuel assembly with the mirror boundary conditions was analysed in the multi group approach. the neutron flux is constant in the assembly with the mirror boundary condition and due to it only eigenvalue was analysed. as it was told before, parcs calculation code offers to use two types of libraries structures (pmaxs and xsec). the eigenvalues were calculated with many different ways: • xsec xsec without downscatter correction, • xsec cor xsec with downscatter correction, 27 pavel suk acta polytechnica ctu proceedings -400 -350 -300 -250 -200 -150 -100 -50 0 50 100 150 200 250 2 4 8 16 23 40 56 r el at iv e d if fe re n ce ( p cm ) group number keff relative difference between scale and parcs xsec corr b1 xsec b1 xsec corr without b1 xsec without b1 figure 8. keff relative deviation with xsec input files • pmaxs pmaxs without downscatter correction, • pmaxs cor pmaxs with downscatter correction. results of the simulations are in table 4. the pmaxs input format brings better results in two group approach, but the large deviations from the reference solution was found in many energy group analysis. the reference solution of the test case geometry was calculated with scale triton sequence keff = 0.86041. it seems that the pmaxs input format is well prepared for two group calculations, which are most common in the full core calculations, but it is absolutely inappropriate for multi groups calculation. the xsec input format seems to be more appropriate for multigroup analysis. there are only small deviations between diffusion solution and appropriate transport solution. the highest deviation with the b1 correction is in case of xsec for two group calculation. this deviation is caused by the fact, that parcs was used in the diffusion mode and the inputs scatter matrix is limited into the downscatter from fast to thermal group. due to that fact, the two group solution needs to be corrected by downscattering via the equation (5). when the downscatter was corrected, the results were similar to those obtained with pmaxs cor. this situation can be seen in both simulations (with b1 and without b1 approach). deviations between the xsec and xsec cor for more group approach are smaller, because for more group calculations, there are total scatter matrix input data. due to it there is no necessary to use correction via equation (5). based on this fact, the xsec without downscatter correction input format is used for other more group calculations. the relative difference of keff between scale and parcs for xsec input format is in figure 8. the convergence problem with the calculation were found in case of higher group number. the calculation did not converge automatically, but after it reached maximum number of iteration, it finished and gave a result without any warning message. the influence of b1 correction decreases with higher number of energy groups. this behaviour begins during the data preparation process by microcode scale triton, because the group structure for the data preparation process was 56 group endf/b-vii.0 library and microcode scale prepares critical spectra in 56 groups. when user wants to calculate xs with b1 leakage correction in 56 groups, the collapsing to the energy groups is independent on the neutron spectra. 4.3. neutron flux spectra comparison the neutron spectra was created to better understand the behaviour in the fuel regions near to the water fuel dummy region. the spectra was calculated with 56 energy groups. there was also the same convergence complications as were described in the previous sections. because of the results of two group calculation, there was analysed only three cases without b1 leakage correction (case4, case5 and case6). the data for scattering were not corrected for downscattering, because the results from section 4.2 showed that the downscattering correction is not necessary. the neutron flux spectra comparison gives the information about the spectra and slowing down in the scale triton and parcs codes. the spectra were analysed separately in fuel regions and in the dum region. 4.3.1. 0.7 % enrichment the test case with 0.7% fuel enrichment represents the sub-critical problem. the neutron spectra in the fuel assembly fa1, respective in the dum region are in figures 9, respective 10. the neutron flux spectra 28 vol. 28/2020 xs data groups with b1 without b1 pmax cor pmax xsec cor xsec pmax cor pmax xsec cor xsec 2 -38 -47 -38 -389 25 25 25 -323 3 -2296 -2307 -93 -106 -1950 -1950 229 228 4 -2168 -2174 40 33 -1952 -1952 231 230 8 -2192 -2195 -7 -10 -1954 -1953 206 207 16 -2110 -2112 77 74 -1955 -1955 205 205 23 -31510 -31512 112 117 -31479 -31479 141 147 40 -31495 -31496 63 68 -31479 -31479 76 83 56 -31985 -31987 -21 -20 -31987 -31987 -21 -20 table 4. keff relative difference between scale and parcs with multigroup approach for 0.7 % enriched fuel assembly with mirror boundary condition 1x10-12 1x10-10 1x10-8 1x10-6 0.0001 0.01 0.01 0.1 1 10 100 1000 10000 100000 1x10 6 1x10 7 fl u x (c m -2 s1 ) energy (ev) neutron spectra in fa1 scale case4 case5 case6 figure 9. neutron flux spectra in the fuel assembly fa1 with 0.7% fuel enrichment 1x10-12 1x10-10 1x10-8 1x10-6 0.0001 0.01 0.01 0.1 1 10 100 1000 10000 100000 1x10 6 1x10 7 fl u x (c m -2 s1 ) energy (ev) neutron spectra in dum scale case4 case5 case6 figure 10. neutron flux spectra in the dum with 0.7% fuel enrichment in fa1 and fa2 are similar and the differences are not visible in the spectra comparison. the eigenvalues and their differences from the reference calculation are in table 5. main result from figures 9 and 10 is that the fuel spectrum in the fuel assemblies is well calculated with the parcs calculation code. differences between the parcs and scale calculation were analysed via equation: ∆φg = φscaleg − φparcsg φscaleg (8) the neutron spectra differences in figures 11, 12 -14 -12 -10 -8 -6 -4 -2 0 2 4 0.01 0.1 1 10 100 1000 10000 100000 1x10 6 1x10 7 fl u x d if fe re n ce ( % ) energy (ev) neutron spectra diferences between scale and parcs in fa1 parcs case4 parcs case5 parcs case6 figure 11. relative difference in spectra between scale and parcs calculation for fa1 for 0.7% fuel enrichment -2 0 2 4 6 8 10 12 0.01 0.1 1 10 100 1000 10000 100000 1x10 6 1x10 7 fl u x d if fe re n ce ( % ) energy (ev) neutron spectra diferences between scale and parcs in fa2 parcs case4 parcs case5 parcs case6 figure 12. relative difference in spectra between scale and parcs calculation for fa2 for 0.7% fuel enrichment and 13 show that the spectrum calculated in the fa1 has the large deviations in the slowing down region and particularly in the high energy neutron region. the fa2 has great deviations only in the high energy neutron region. in the fuel assemblies, the relative deviations are only up to 12% in the absolute values, but the deviations in the dum region are up to 225%. that means that the data prepared for the dum region are not prepared well and the behaviour of neutrons near the dum region is not described correctly. 29 pavel suk acta polytechnica ctu proceedings -250 -200 -150 -100 -50 0 50 100 0.01 0.1 1 10 100 1000 10000 100000 1x10 6 1x10 7 fl u x d if fe re n ce ( % ) energy (ev) neutron spectra diferences between scale and parcs in ref parcs case4 parcs case5 parcs case6 figure 13. relative difference in spectra between scale and parcs calculation for dum for 0.7% fuel enrichment 4.3.2. 4.3 % enrichment fuel pins with 4.3% enrichment represent the supercritical problem, which, based on the information obtained during the two group study, can brings more deviated results. the eigenvalue differences are in table 5. the neutron spectra were similar to the neutron spectra in the lower enrichment case with little bit higher discrepancies. mode ∆keff 0.7% (pcm) ∆keff 4.3% (pcm) case4 120 923 case5 24 783 case6 61 837 table 5. keff relative difference between scale and parcs with 56 groups calculation 4.4. sp3 multigroup solution the parcs calculation code is able to calculate simplified transport approach (sp3) instead of the diffusion approach. the nemmg solver with the nspn 3 was used to calculate the test geometry. the only appropriate approach of case4, case5 and case6 were calculated. the results of the eigenvalue for both enrichment cases are in table 6. mode ∆keff 0.7% (pcm) ∆keff 4.3% (pcm) case4 -52 732 case5 -148 590 case6 -111 646 table 6. keff relative difference between scale and parcs with 56 groups calculation and sp3 the neutron spectra for sp3 calculations are very similar to the neutron spectra calculated with diffusion approach. the main difference is in the higher energies in the fa2 fuel assembly with 0.7% enrichment. the neutron spectra are more accurate in the higher energies in comparison with the diffusion approach for higher enriched fuel. the eigenvalue deviations increased in comparison to the diffusion approach in lower enrichment case, but deviation decreased with higher enrichment case. 5. conclusion various different approaches were analysed in this paper. two different enrichments were studied. results of the two group calculation showed that in case of lower enrichment, the data prepared by the simulation of the fa1 fuel assembly in the infinite lattice and fa2 and dum in the real geometry with b1 approximation (case3) well reproduce the eigenvalue. the relative flux differences were also calculated with the smallest difference in the case3. on the other hand, the cases, which were calculated without b1 approximations gave more-times worse results in case of lower enrichment. the results were opposite in the case of higher enrichment than in the lower enrichment calculation. the best agreement were obtained during the calculation without b1 approximation and the same data preparation scheme (case6). the eigenvalue was bellow 100 pcm from reference calculation, but the fast group neutron fluxes were calculated with higher difference (up to 10%) in each calculation without b1 approximation. the upper results are consistent with b1 correction. in the case of lower enrichment, the situation corresponds with the physical solution. the whole geometry is sub-critical and b1 approximation decreases the neutron leakage to reach criticality during the data preparation process. in case of higher enrichment, the whole geometry is super-critical, but there is part which is "locally" sub-critical dum. the b1 approximation increases neutron leakage from the system to reach the criticality, but this does not correspond with the physical solution in the dum part. the neutron spectra is changed there and due to that the b1 correction cannot brings better results in the case of higher enrichment. the main objective of the spectra study is that pmaxs files, which parcs use newly, are not able to calculate core with multigroup approach. the eigenvalue differences showed that there is any problem with data preparation for that calculation. the results of calculations with xsec file format bring reasonably results of eigenvalue despite the convergence problems. based on the results of multigroup study, the neutron spectra in multigroup approach were calculated for test cases without b1 approximation. calculations without b1 approximation were chosen because the final geometry was not critical and there are still unsolved problems withb1 approximation. the paper [7] informs that the b1 calculation is not rigorous and due to the reaching critical spectra by the changing absorption in each energy group, the result are not representative for lwr calculations, where the criticality is reached by the absorption of only thermal neutrons. the neutron spectra in the fuel assemblies were determined with a good agreement with the reference 30 vol. 28/2020 xs data calculation, but the spectra in the dum was determined with larger discrepancies. it is well known, that the diffusion theory is not fulfilled in the water fuel dummy region and the discrepancies are mostly based on this fact. there are also methods, which can bring better results, for instance the influence of neutron scatter anisotropy in the dum. the different approach of the dum preparation data will be study in the future. 6. acknowledgement this work was supported by the project no. cz.02.1.01/0.0/0.0/16_013/0001790 supported by operational programme research, development and education co-financed from european structural and investment funds and from state budget of the czech republic and it was supported by the grant agency of the czech technical university in prague, grant no. sgs19/114/ohk4/2t/14. list of symbols φi(e) neutron flux of region i and energy e [cm−2s−1] φi,g neutron flux of region i and energy group g [cm−2s−1] σi,g microscopic cross section of region i and group g [cm2] σi(e) microscopic cross section of region i and energy e [cm2] σi,g macroscopic cross section of region i and group g [cm−1] ni atomic density of region i [cm−3] eg, min minimal energy of boundary related to the energy group g [ev] eg, max maximal energy of boundary related to the energy group g [ev] n number of regions in the test case [–] σt,g total macroscopic cross section for energy group g [cm−1] b buckling factor [cm−1] jg neutron current in group g [cm−2s−1] χg neutron spectra in group g [–] σf,h fission macroscopic cross section [cm−1] σ0s,h→g scattering macroscopic cross section [cm −1] keff,x eigenvalue calculated with mode x [–] ∆keff difference of the eigenvalue [pcm] ∆φg relative difference of φg for the whole test case [%] references [1] s. choi, k. s. smith, h. kim, et al. on the diffusion coefficient calculation in two-step light water reactor core analysis. journal of nuclear science and technology 54(6):705–715, 2017. doi:10.1080/00223131.2017.1299648. [2] a. dall’osso. a spatial rehomogenization method in nodal calculations. annals of nuclear energy 33:869–877, 2006. doi:10.1016/j.anucene.2006.05.008. [3] p. suk. advanced homogenization methods for pressurized water reactors. acta polytechnica ctu proceedings 19:14, 2018. doi:10.14311/app.2018.19.0014. [4] p. reus. neutron physics. 2008. isbn: 2759800415. [5] k. i usheva, s. kutsen, a. a khruschinsky, l. babicheu. generation of xs library for the reflector of vver reactor core using monte carlo code serpent. journal of physics: conference series 781:012029, 2017. doi:10.1088/1742-6596/781/1/012029. [6] t. clerc, a. hébert, h. leroyer, et al. an advanced computational scheme for the optimization of 2d radial reflector calculations in pressurized water reactors. nuclear engineering and design 273:560 – 575, 2014. doi:10.1016/j.nucengdes.2014.03.044. [7] k. s. smith. nodal diffusion methods and lattice physics data in lwr analyses: understanding numerous subtle details. progress in nuclear energy 101:360 – 369, 2017. special issue on the physics of reactors international conference physor 2016: unifying theory and experiments in the 21st century, doi:10.1016/j.pnucene.2017.06.013. [8] t. downar, y. xu, v. seker. parcs v3.0 u.s. nrc core neutronics simulator theory manual. res/u.s. nrc, rockville, md, 2012. [9] b. t. rearden, m. jessee. scale code system. ornl/tm-2005/39, 2016. [10] j. leppänen. serpent a continous-energy monte carlo reactor physics burnup calculation code. vtt technical research centre of finland, 2015. [11] purdue advanced reactor core simulator. purdue university accessed: 2019-05-14. https://engineering.purdue.edu/parcs. [12] b. t. rearden, m. jessee. scale code system. ornl/tm-2005/39, 2018. [13] a. ward, t. downar, y. xu. genpmaxs v6.2 code for generation the parcs cross section interface file pmaxs. res/u.s. nrc, rockville, md, 2016. http: //nuram.engin.umich.edu/software/genpmaxs/. 31 https://doi.org/10.1080/00223131.2017.1299648 https://doi.org/10.1016/j.anucene.2006.05.008 https://doi.org/10.14311/app.2018.19.0014 https://doi.org/10.1088/1742-6596/781/1/012029 https://doi.org/10.1016/j.nucengdes.2014.03.044 https://doi.org/10.1016/j.pnucene.2017.06.013 https://engineering.purdue.edu/parcs http://nuram.engin.umich.edu/software/genpmaxs/ http://nuram.engin.umich.edu/software/genpmaxs/ acta polytechnica ctu proceedings 28(0):1–9, 2020 1 data preparation process and reflector calculation 1.1 macroscopic cross section preparation 1.2 b1 correction 2 calculation codes 3 test case 4 calculation 4.1 two group approach 4.1.1 0.7 % enrichment 4.1.2 4.3 % enrichment 4.1.3 discussion 4.2 multigroup approach 4.3 neutron flux spectra comparison 4.3.1 0.7 % enrichment 4.3.2 4.3 % enrichment 4.4 sp3 multigroup solution 5 conclusion 6 acknowledgement list of symbols references https://doi.org/10.14311/app.2022.33.0125 acta polytechnica ctu proceedings 33:125–132, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague high thermal conductivity concrete for energy piles luca di girolamoa, ∗, gigliola ausiellob, gianpiero russoa, gabriella maronea a university of naples federico ii, polytechnique and basic sciences school, department of civil, building and enviromental engineering, via claudio 21, 80125, napoli, italy b university of naples federico ii, polytechnique and basic sciences school, department of civil, building and enviromental engineering, p.le tecchio 80, 80125, napoli, italy ∗ corresponding author: luca.digirolamo@unina.it abstract. research is increasingly focusing on thermal properties of concrete with the aim of reducing the heat exchange between buildings and environment. on the other hand, concretes with high thermal conductivity could have interesting applications in the field of thermo-active ground structures as geothermal energy piles (geps). this kind of foundations represent an environmentally friendly technology that allows exploiting the heat of the shallow earth surface to supply renewable energy for the air conditioning of buildings. geps are needed for structural and geotechnical reasons and allow recovering the installation costs connected to vertical boreholes. concrete drilled or driven piles are equipped with a primary circuit (pc) of high-density polyethylene plastic pipes attached to the reinforcement cages. thermal energy is extracted from or injected into the ground thought a carrier fluid that flows into the pipes of the pc. to improve the heat exchange between the pile and soil the thermal properties of the concrete should be considered as design parameters. concrete thermal conductivity, contrary to what happens for the buildings, should be increased to optimise the thermal performance of the geps. different solutions that modify the mix design of concrete are proposed to the aim of increasing the thermal performance of geps. keywords: energy piles, high conductivity concrete, sustainability. 1. introduction energy demand increasing is linked to the rapid economic development that is taking place worldwide. however, the energy requirement cannot continue to be satisfied using fossil energy sources which are responsible for climate changes and environmental pollution [1]. the use of renewable energy source is promoted by european parliament directive 2010/31/ue on energy performances of buildings [2]. geothermal energy is one out of the possible and most easily available renewable energy sources. in fact, below the ground surface, at depths greater than 10 m b.g.l, temperature of ground is nearly constant throughout the year. using this relatively constant temperature, the air conditioning of buildings, both in the heating phase and in the cooling phase, it is more efficient than other heating or cooling systems [3]. in the early 80’s geothermal energy was initially extracted, in some parts of central europe like austria and switzerland, from deep foundation elements like piles. in this technology, differently from conventional ground heat exchangers made by one or more u-shaped plastic absorber pile inserted in aăborehole, geps use thermal conductivity and thermal storage capacity of concrete because high-density polyethylene plastic pipes are installed directly in piles structures before concrete casting. thermal energy is extracted from or injected into the ground via a heat carrier fluid that flows into the pipes of the pc [4]. since geps combine both the structural and the energetic functions in a single solution, they are known to be cost effective [5]. in gep’s technology, concrete can play a very important role. usually, the intrinsic characteristics of this material, in relation to its traditional applications for super-structures, do not allow to have high performances from the energetic point of view because its relatively high thermal conductivity is the main cause of heat dispersion inside the buildings. for geothermal applications, concrete can gain a redemption opportunity. its high thermal conductivity can be further increased, increasing at the same time the efficiency of the geothermal plant and without dramatic increases of the stress level within the structural sections of the piles. on this last aspect of course, a generalisation is not opportune and any case should be carefully designed and checked. in this paper, the role of concrete in geps technology is analysed. the target is to analyse the effects of thermal conductivity increasing of concrete on energy and mechanical performances of geps. after an introduction about thermal properties of concrete and a review of literature about ways to enhance its ther125 https://doi.org/10.14311/app.2022.33.0125 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en l. di girolamo, g. ausiello, g. russo, g. marone acta polytechnica ctu proceedings figure 1. description of mechanism of heat transfer from primary circuit of energy pile to ground. mal conductivity, the evaluation of stresses in a geps structural was carried out through a finite element (fe) code software. 2. heat transfer in geothermal concrete energy piles heat transfer mechanism from the primary circuit of geps to soil (and vice versa) during heating or cooling of built structures is a complex mechanism that involves all the underground components of geothermal system (geps). primary circuit consists of heat exchanging hdpe loops incorporated within piles foundation, heat exchangers fluid that flows inside pipes loops and concrete of the pile. between the fluid circulating into the pipes inside the pile and the surrounding soil there is a thermal gradient due to differences of temperature and thermal properties of the fluid and the soil [6]. heat transfer in a concrete energy pile occurs by heat convection between carrier fluid and wall of pipes and heat conduction between pipes’ wall and concrete of pile and between concrete and soil [7]. usually, for heat transfer in geps convection and radiation mechanisms are negligible [8]. according to loveridge et al. [9] the larger dimensions of the diameter of the geps compared to traditional geothermal wells ensure that it cannot be considered a thermal steady state. in fact, since the concrete around the pipes can take several days to reach the steady state, the difference in the average temperature between the heat exchanger fluid and the average soil temperature on the edge of the exchanger cannot be considered constant. for this reason, thermal resistance of concrete could become a fundamental parameter to improve thermal performances of geps. the heat transfer between the primary circuit fluid and the soil is given as [6]: q = t1 − t5 rt ot (1) where t1 and t5 are fluid and soil temperatures, respectively, and rt ot is the total thermal resistance transfer, given as: rt ot = rf luid + rpipe + rconcrete + rground (2) the resistance of the following fluid is given as: rf luid = 1 2nπrih (3) where ri is the internal pipe radius, n is the number of pipes, and h is the convective heat transfer coefficient. the pipe thermal resistance is given as: rpipe = ln (r0/ri) 2nπkp (4) where r0 is the pipe outer radius and kp is the thermal conductivity of the pipe material. regarding thermal resistance of the concrete annular section, its steady state value could be calculated with the equation of the thermal resistance of a cylinder, but an assumption must be made for the effective inner radius of that cylinder ref f . the concrete thermal resistance using the equivalent diameter approach is given as: rconcrete = ln (rb/ref f ) 2πkc (5) where rb is pile radius, kc is concrete thermal conductivity, and ref f is the effective radius ref f = r0 √ n (6) the equivalent cylinder approach does not consider actual positioning of pipes. the mechanism is explained in figure 1. 126 vol. 33/2022 high thermal conductivity concrete for energy piles figure 2. cooling (a) and heating (b) energy usage of a building in chicago compared to different concrete thermal conductivities (image reprinted from [6] and [10]). 3. heat transfer enhancement: thermal conductivity of concrete equation 2 and the followings show that thermal conductivity k and/or convective heat transfer coefficient h increasing correspond to thermal resistance decreasing. in this way, the heat transfer between the primary circuit fluid and the soil will increase. according to kwang et al. [10], increasing foundation thermal conductivity reduces the heating and cooling energy demands of a building (figure 2). therefore, to increase heat exchange between energy piles and soil, designers could work on thermal properties of thermal fluid, thermal properties of pipes, by optimizing pile dimensions and geometrical arrangement of pipes or by improving thermal properties of concrete modifying its mix design. this study focused on this later aspect. conduction is the most important mechanism of heat transfer for concrete. thermal and mechanical properties of concrete are influenced by its porosity [11]. according to asadi et al. [12], moisture, the specimen’s condition and aggregate volume fraction appeared to be the main factors influencing the thermal conductivity of concrete. however, the most effective factors on the thermal conductivity are the water/cement ratio and type of admixture. the analysis of the variables that influence thermal conductivity, although in many studies is aimed at its reduction, is oriented towards a useful definition of a relationship between the density of concrete and its conductivity. anyway, because the target is the realization of geps, any modification of mix design should not alter piles’ resistance. this kind of foundations in comparison to the ordinary piles are subjected to additional loading induced by thermal variations. additional strains and stresses induced by the thermo mechanical coupled loadings occur. from this point of view the improvement of both thermal and mechanical properties could be an interesting point. this aim could be achieved developing mix designs which can boost compactness and density at the same time. according to kim [13] thermal conductivities of concrete mixtures were revealed independent of curing age. in relation to the components of the mixture and the relationships between them, kim et al. (2003) show interesting results. first, the larger the amount of coarse aggregate in a concrete mix the higher the thermal conductivity of the mixture because of the well-known higher thermal conductivity value of aggregate than other constituents of the concrete mix. furthermore, the nature of the aggregate also appears to have an influence on conductivity [14]. the siliceous aggregate, compared to the carbonate one, having a more regular molecular structure (crystallinity), has greater thermal conductivity. even with an increase in the fine aggregate, an increase in thermal conductivity is obtained [13]. low values of w/c ratio lead to higher conductivity of the concrete mix. in fact, the thermal conductivity of the cement is higher than that of the water and above all the water is responsible for the formation of the voids that are created after its natural evaporation. so, a lower w/c ratio corresponds to a lower porosity of the concrete. thermal conductivity decreases linearly with temperature but this dependence is spread over large temperature ranges. as a matter of fact, in the temperature range in which the geps work the dependence of the conductivity from the temperature it is nearly negligible. finally, thermal conductivity increases when voids of concrete are filled with water instead than air. zhang et al. [15] shows that under saturated conditions thermal conductivity increases by at least 50%. other studies [16–19] provide reports to parameterize the increases in the k value as a function of the hu127 l. di girolamo, g. ausiello, g. russo, g. marone acta polytechnica ctu proceedings midity and of the weight of the concrete both due to the absorption of water. therefore, pile foundations made by concrete directly cast in place in soils below the groundwater table can boast greater performance as geothermal heat exchanger. starting from these assumptions, it is possible to identify some precautions that mix design should have to optimize both the structural and the energy performance. the compactness to be maximized is certainly the right solution, which can be obviously obtained following different procedures. high strength and significant compactness also make the cement matrix physically impenetrable to the aggressions of the environment, the microporosity being reduced too. both are a guarantee of durability, a fundamental aspect of the performance for any concrete structural object and this is especially true for a foundation pile and even more for an energy foundation pile. in the literature, several authors [20–24] have carried out lcc and lca evaluations of geothermal pump systems, equipped with energy piles also. they have also compared them with more conventional heating and cooling systems. these studies have shown a reduction of environmental impact of ghps than conventional plants. in economic terms, according to morrone et al, the npv (net present value) varies from about €12.000 to about €33.000 while the payback period varies from 4 to 11 years for geps installations depending on the climate zone considered in italy [24]. however, other considerations and insights on lcc and lca deserve a separate study. 4. guidelines for a mix design with high thermal conductivity for piles in new concrete mixes, fillers are increasingly used with the main function of enhancing the workability in the wet state. it is known, however, that the performance changes are mainly characterized by an increase in the resistance, since the presence of the filler reduces the microporosity due to acompactness increasing. calcareous filler, with its physical action, only increases the compactness, while the pozzolanic filler, with its chemical action, also improves the resistance, involving lime in the formation of a second family of hydrate compounds [25]. kim [13] and demirboğa [26] test the effects of some pozzolanic fillers in partial replacement of cement and detect reductions in thermal conductivity. in fact, compared to the large and fine aggregate, the main function of the filler is that of a micrometric aggregate that can reduce microporosity. therefore, inserted in the mix design as a new component, it can only increase the compactness at the micrometric scale and, consequently, increase density and mechanical resistance. this creates the conditions for greater thermal conductivity. another correction of the mixture is the use of the new superplasticizing additives, also called water reducers, which allow to keep with the same degree of workability a low value of the w/c ratio. this suggestion among many of the guidelines on mix composition is the most important because thermal conductivity is inversely proportional to the w/c ratio. as a result, the viscosity is reduced without causing the segregation of the concrete. in addition, with their deflocculating action, they limit the formation of clots of the cement granules, which would form when dispersed in the water. fillers and superplasticizers are two additional components that enter the new self-compacting concretes, in response to construction site management needs. with rather low w/c ratios and the consequence of reducing microporosity, self-compacting concrete (scc) [25] achieves high resistance values. a "technology in technology" of scc concrete is that of continuous flight auger (cfa) piles. it is a technique to install replacement piles combining and axial thrust and an action of a torque on a continuous flight auger inserted in soil. the scc could be pumped to the down hollow stem of the auger thanks to the exceptional workability characteristics. while scc is pumped, without rotation the auger is pulled out. in this way, temporary supported are not necessary. once the concrete has been cast in place pre-assembled reinforcement cage, is easily inserted into wet concrete. in the case of geps also geothermal pipes are fixed on the pre-assembled reinforcement cage. it could be added that the scc casting performed from bottom allows the mixture to incorporate less air and therefore to reduce even more the microporosity. the possibility of execution in soils of any nature, even in the presence of groundwater, widens the application horizon, combining greater conductivity and mechanical resistance with easy execution on site, without cost increases [25, 27]. asadi [12] relates the thermal conductivity of scc concrete with the increase in temperature, highlighting intervals with increases and decreases starting from the room temperature value and reports the results of some studies done by khaliq and kodur [28] suggesting the following relationship between thermal conductivity and temperature for a range of temperature of concrete between 20◦c and 400 ◦c (equation 7). k = 3.12 − 0.0045t (7) as can be seen from equation 7, the increase in conductivity is inversely proportional to the temperature of concrete. considering that the range of temperatures at which the geps work is very close to the lower limit of the range proposed by the authors, it follows that the conductivity values of the concrete remain quite high. the use of steel fibers seems another interesting guideline, considering that the thermal conductivity of steel is approximately 50 times greater than 128 vol. 33/2022 high thermal conductivity concrete for energy piles that of concrete and that the density of steel is approximately 3ătimes that of concrete. adeyanju and manoha [29] show that the fiber reinforced concrete (frc) with steel fibers reaches high levels of thermal conductivity, of about 2,0-2,5 w/m◦c compared to 1,01,3ăw/m◦c of a concrete without fibers, with increases well above 100%. this increase is also accompanied by a higher density, once again signing a direct proportionality of the conductivity with both compactness and density. nagy, nehme and szagri [30], comparing various investigations on frcs with steel fibers, they find out values of thermal conductivity ranging from 2,0 to w/m◦c 3,2 w/m◦c. on the other hand, khaliq, kodur, experiment with composite details. these are fiber-reinforced sccs with thermal conductivity values ranging from 3,0 w/m◦c to 3,5 w/m◦c. obviously, the size and quantity of the fibers also influence the formation of the voids. indeed, nagy, nehme and szagri [30] find out that by increasing the quantity of steel fibers the thermal conductivity starts to decrease. this phenomenon is due to the formation of voids for effect of the greater quantity of fibers. lie and kodur [14] also report an interesting relationship between thermal conductivity and temperature in a concrete in which the effects of steel fibers and the nature of the aggregate are added. equation 8 is related to siliceous aggregate for a range of concrete temperature between 0 ◦c and 200 ◦c, while equation 9 is related to carbonates aggregate for a range of concrete temperature between 0 ◦c and 500 ◦c. k = 3.22 − 0.007t (8) k = 2.00 − 0.001775t (9) the comparisons show the order of magnitude of the thermal conductivity k which, at temperatures close to those of the soil, presents high values when the compactness is high and there is the right amount of steel fibers. furthermore, in all these mixtures with steel fiber reinforcement, characterized by high mechanical properties, the action of the fibers does not induce any improvement in compressive strength, but only in tensile one. furthermore, metal fibers have little influence on the heat capacity at low temperatures [14] even for high quantities of fibers [30]. the integration with microcapsules in pcm is another possible guideline. asadi [12] found that the thermal conductivity of cementitious materials with pcm is lower than conventional cementitious materials and several authors have parameterized this decreasing [31, 32]. the encapsulation system is crucial because the use of small quantities of microcapsules could reduce the microporosity. pcm microcapsules could increase thermal capacity without having too strong effects on the reduction of thermal conductivity, improving the storage capacity of the concrete dampening indirectly the already modest temperature fluctuations affecting the soil supporting the steady-state hypothesis. as shown in the section there are a lot of potential paths to enhance the efficiency of concrete piles as heat exchangers optimizing the heat transmission mechanism. 5. effect of concrete thermal conductivityăon pile-soil thermo-mechanical interaction heat transfer enhancement in concrete improves the heat transfer process among the heat exchanger piles and the soil. several studies report the beneficial effects in terms of energy performance of geothermal systems, but the conductivity improvement should be considered also in terms of load carrying capacity of the pile foundation that should however not be compromised [5]. if the heat transfer phenomena inside the pile are modified, additional investigations about pile-soil thermo-mechanical interaction are needed. herein the effects of concrete thermal conductivity improvement on the short term thermo-mechanical behaviour of a single energy pile are briefly presented. four different cases are compared. case 0 where concrete thermal conductivity is assumed 2,4 w/mk, case 1, case 2 and case 3 where thermal conductivity is assumed 2,88 w/mk, 3,6 w/mk and 4,8 w/mk respectively. a cfa pile of 13 m length with 0,60 m diameter, embedded in a multi-layered pyroclastic deposit with a socket in yellow tuff 3,00 m deep is modelled. the pile is the prototype used recently in the foundation project of a new trading centre in city of napoli [33]. transient fully coupled analyses are carried out with the fe software plaxis 2d. one day of heating, characterized by a maximum temperature variation of 13 ◦c, and one day of cooling, characterized by a maximum temperature variation of 17 ◦c, are simulated [34]. to reproduce realistic operational conditions, thermal loadings are combined to the live mechanical top load. the mechanical load, assumed as the 40% of the pile bearing capacity, is thus applied and kept constant during thermal loadings fluctuations. the effects of the thermal conductivity on pile-soil thermo mechanical interaction are analysed in terms of axial loading along the pile shaft. all the thermo-mechanical cases presented (case 0, case 1, case 2 and case 3) are compared to the purely mechanical case (mechanical) where only the live mechanical top load, (i.e. 2400 kn), is applied to the pile. the axial load along pile shaft is reported during and at the end of the cooling phase (respectively labelled c for cooling and ec for end cooling) and during and at the end of the heating phase (h for heating and eh for end of heating) for each thermomechanical case (figure 3). from figure 3 it could be 129 l. di girolamo, g. ausiello, g. russo, g. marone acta polytechnica ctu proceedings figure 3. c axial loads along the pile shaft for different thermal conductivity of concrete. observed that an improvement of concrete thermal conductivity, for a daily thermal cycle, does not determine significant effect on the thermal axial loading both for cooling and heating modes. for case 1 and case 2 the maximum thermal axial load increasing during heating and decreasing during cooling are coincident and equal to 390 kn and 320 kn respectively. in case 3 the maximum thermal axial load increasing during heating and decreasing during cooling are almost 380 kn and 330 kn respectively. in terms of axial load, increments of 20%,50% and 100% of the thermal conductivity of concrete determine comparable effects. comparing case 0 to case 3, corresponding to minimum and maximum thermal conductivity respectively, a difference of almost 4% of the thermal axial load is observed. the enhance of concrete thermal conductivity, for a short time analysis, seems not to determine significant axial thermal load variations. for long term analyses should also be considered that an enhance of thermal conductivity could determine a beneficial effect in terms of temperature distribution on the cross section and therefore a more uniform axial stress distribution. as a matter of fact, normal stresses induced by mechanical and thermal loadings are not uniform in the cross section. the trend of the axial stress is characterised by sudden variations close to the primary circuit position. an improvement of thermal properties of concrete could mitigate this effect and could be beneficial even in terms of durability of concrete. 6. conclusions in this paper the thermal conductivity of concrete and the use of high conductivity concretes for special applications like geps was analysed. after an introduction about thermal properties of concrete and a review of literature about ways to enhance its thermal conductivity, guidelines for a mix design with high thermal conductivity was laid out and simulations through fe analyses about the potential effects of the thermal improvement on mechanical behaviour were reported. the axial stress induced by live loadings combined to heating or cooling was evaluated for a potential energy cfa pile characterised by different thermal conductivities. for a concrete with high thermal conductivity low w/c ratio is necessary to reduce the voids inside the cement matrix. for this reason, the use of fillers, superplasticizers or both at the same time is desirable to achieve the goal. the addition of steel fibers within the mix design of concrete is another possible expedient. in addition to improving the mechanical capabilities of concrete, they also increase its thermal conductivity with negligible influence on thermal capacity at low temperatures. the effect of microcapsules with pcm should not affect conductivity very much but an increase in the thermal capacity is typically expected. the simulations carried out with fe software for a prototype pile embedded in a layered soil profile similar to a real foundation project in the eastern area of the city of napoli has shown that the change in the thermal properties of the concreted is not producing large effects on the structural performance. computed axial compression loading changes for the thermal induced strains is in a range of ±15% compared to the axial loadings obtained by the head load corresponding to a typical 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[34] g. marone, l. di girolamo, g. russo. studio parametrico di un palo singolo sottoposto a carichi termomeccanici, iagig 2019, 2019. 132 https://doi.org/10.1016/j.cemconres.2011.06.012 https://doi.org/10.4236/jmmce.2011.1015111 https://doi.org/10.1016/j.egypro.2015.11.616 https://doi.org/10.1016/j.conbuildmat.2016.12.185 https://doi.org/10.1016/j.conbuildmat.2012.04.042 https://doi.org/10.1007/s11440-012-0177-4 https://doi.org/10.14311/app.2022.33.0092 acta polytechnica ctu proceedings 33:92–97, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague strength and durability characteristics of self compacting concrete (scc) with recycled aggregate and manufactured sand sashidhar chundupallea, ∗, nirmala seeria, sudarsana reddy veerab, venkata ramana nelluric a jawaharlal nehru technological university anantapur, department of civil engineering, 28-3-322-2, sarada nagar ii cross, 515002 anantapur, india b priyadarshini engineering college, department of civil engineering, lakshmi nagar, ramachandrapuram-pudi rd, ramachandrapuram, 517561, tirupati, india c university b.d.t college of engineering, department of civil engineering, university bdt college of engineering, 577004 davanagere, india ∗ corresponding author: sashidhar.civil@jntua.ac.in abstract. the investigations on self compacting concrete (scc) with recycled coarse aggregate and manufactured sand (m-sand) were performed in this current study. the scc mixtures were produced with the recycled coarse aggregate of proportions of 0 to 100% with step increment of 25% with a ratio of water binder at 0.36. in addition, the mixtures were also prepared by mixing it with m-sand. the feasibility of utilizing the recycled coarse aggregate in scc was evaluated through strength and durability studies. the obtained results demonstrated that the scc produced from these materials can be effectively recommended for their usages in concrete industries. keywords: m-sand, mechanical properties, recycled coarse aggregate, self compacting concrete. 1. introduction self compacting concrete (scc) show superior rheological properties as compared to conventional concrete. these properties in scc are generally achieved due to higher ratio of cement paste to aggregate. accordingly, the scc was developed to overcome the problems associated with the labor quality and the noise and dust problems as well. the mechanical characteristics such as tensile and compressive strength of concrete evaluated with recycled aggregate will be lower than the typical concrete. however, this can be overcome by using superplasticizer [1, 2]. the studies conducted on the fresh and hardened scc using recycled coarse and fine aggregate confirmed that both the coarse as well as fine recycled aggregate can be effectively used for the production of scc [3, 4]. accordingly, in this study, an equal consistency was observed in scc mixtures when it is prepared with the recycled coarse aggregate with 0, 50 and 100% proportions as a substitute for the natural coarse aggregate [5]. regarding the fine aggregate, the manufacturing sand (commonly known as m-sand) can be a suitable alternative for river sand in concrete. the incorporation of m-sand contributes to increase in paste volume, which eventually helps for scc for its improved properties. it means that as compared to the river sand, the m-sand produces relatively high paste content, which is beneficial for scc [6]. the utilization of recycled coarse aggregate of 0 to 40% proportions (with a step increment by 10%) in scc as a substitute for the natural coarse aggregate showed relatively decreased mechanical strength as compared to those made with natural aggregate [7, 8]. on the other hand, it was reported that the utilization of fine and recycled coarse aggregate enhances the rheological and fresh properties of scc [9]. few authors [10, 11] have used the crushed red bricks and crushed ceramics materials as a replacement of coarse aggregate. however, the scc made by m-sand and fly ash contributed to raise of the strength characteristics of scc [12]. from the above discussion, it is observed that only a very few studies were carried out with amalgamation of coarse and fine aggregate as suitable replacements in scc. hence, the current experimental study is focused on enhancing the strength characteristics and durability of scc through the substitution of both coarse and fine aggregate, where these aggregate were substituted with recycled coarse aggregate and m-sand, respectively at the proportions of 0, 25, 50, 75 and 100%. 2. experiments and methods 2.1. materials in this current study, the conventional portland cement grade 53 with specific gravity (sg) of 3.05 and fineness value of 310 m2/kg, was used. similarly, the class f’ fly ash was found to be as per the astm c 618 standard with a relative density of 2.13 and fineness value of 360 m2/kg. the chemical ingredients of 92 https://doi.org/10.14311/app.2022.33.0092 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 analysis of concrete and cement epd figure 1. photographic image of the experimental setup for rapid chloride permeability test. the cementitious materials are tabulated in table 1. natural granite was used as a coarse aggregate with a max. size of 1.2 cm with sg around 2.66 and the water absorption value of 0.53. however, the local demolished wastes were also used as the recycled aggregate, which was with a sg around 2.52 and water absorption value of 5.06. the local river sand with sg around 2.58 and water absorption value of 0.95 was employed as the fine aggregate components. the sg and water absorption value of m-sand used in the study are 2.46 and 1.20, respectively. a chemical based superplasticizer named polycarbolic ether was also used to achieve a good workability. chemical present cement fly ash value (%) sio2 19.79 65.6 al2o3 5.67 28.1 fe2o3 4.68 2.9 cao 61.81 0.95 mgo 0.84 1.55 so3 2.48 0.21 chloride 0.003 lime saturation factor 0.92 table 1. chemical compositions of scc. 2.2. mix proportions scc mixtures at proportions of 0 to 100% replacements (with step increment by 25%) with the recycled aggregate and m-sand were prepared, while keeping the ration of water binder (w b) constantly at 0.36. throughout this work, cement and fly ash was taken to be 351 and 150 kg/m3, respectively. 2.3. test methods in this current work, the scc prepared using the aggregate of recycled coarse and m-sand was tested as per the efnarc and is guidelines [13]. 2.3.1. split tensile and compressive strength test the split tensile and compressive strength characteristics of various mixtures were evaluated on the test setup with capacity of 2000kn compression testing machine with a least count of 1kn. 2.3.2. rapid chloride permeability test (rcpt) the rcp tests were conducted on the different concrete mixtures in order to investigate the developed concrete’ ability to resist the penetration of chloride ions. for this purpose, the samples were cut into cylindrical specimens and subjected to impressed voltage of 60v between the anode and cathode using a dc power source as shown in fig 1. the response current was monitored without any interruption for the period of 6 h at 30 min of interval time. 2.3.3. acid resistance test the resistance against the acid attack on scc specimens was conducted in accordance with the astm c 267 01. after the process of moist curing for 28 days, the weight of each specimen has been taken and then the specimens were immersed in 3% hydrochloric (hcl) acid solution. during this test, the changes in weight as well as the compressive strength characteristics of specimens were evaluated after 28 days and 56 days of immersion of the specimens in 3% hcl acid solution. 3. results and discussion 3.1. compressive strength analysis 3.1.1. compressive characteristic strength of rca in nca (natural coarse aggregate) concrete mix with 0% m-sand the compressive characteristic strength results of the developed various concrete mixtures are displayed in 93 s. chundupalle, n. seeri, s. r. veera, v. r. nelluri acta polytechnica ctu proceedings s. no mix type average strength (mpa) m0/days m25/days m50/days m75/days m100/days 7 28 7 28 7 28 7 28 7 28 1 r0 29.77 42.66 28.88 41.77 28.1 40.44 27.42 40.00 27.07 39.11 2 r25 28.27 38.88 27.11 38.66 26.66 38.00 26.22 37.55 25.86 36.6 3 r50 26.76 36.24 25.55 35.92 25.33 35.57 24.88 35.14 24.44 34.52 4 r75 25.00 33.77 23.88 33.33 23.53 32.88 23.35 32.44 23.00 32.00 5 r100 23.77 31.58 22.66 31.11 22.00 30.44 21.55 30.11 21.33 29.77 r0 first letter stands for rca and the suffix number indicates the % used m0 first letter stands for m-sand and the suffix number indicates the % used table 2. estimated compressive strength of various concrete mixtures. table 2. from the obtained results, it can clearly be noticed that the increasing concentration of recycled aggregate concrete considerably decreases the effective compressive strength continuously, which is also well agreed with the literature reports. further, it can also be noted that the compressive strength of 100% nca is maximum and the compressive strength decreases with increasing recycled coarse aggregate at both 7 and 28 days. this might be due to the poor quality of the recycled aggregate as compared to the natural aggregate. the percentage decrease in the compressive strength for 7 days sample with reference to the basic mixture is found to be 7.45, 13.43, 19.38 and 22.37% for 25, 50, 75 and 100% replacement of recycled coarse, respectively. similarly, the percentage decrease in the compressive strength for 28 days sample with reference to basic mix is estimated to be 8.86, 14.58, 20.83 and 23.62% for 25, 50, 75 and 100% replacement of recycled coarse, respectively. 3.1.2. compressive strength of m-sand substitution in fine aggregate with 0% substitution of recycled coarse the compressive investigation results of replacing the m-sand in fine aggregate are given in table 2. in this work, as aforementioned, the fine aggregate has been replaced by m-sand at different percentages such as 0, 25, 50, 75 and 100% and the percentage amount of recycled aggregate was taken to be 0%. from the test results of the obtained compressive strength, it is noticed that the increasing substitution of m-sand leads to a continuous decrement in the eventual compressive strength of the concrete mix. accordingly, the strength property of 28 days immersed 100% msand scc is estimated to be 29.77 n/mm2 and on the other hand, it can be seen that the compressive strength of cubes is decreasing with the increasing m-sand substitution in fine aggregate. this may be because of the reason that the crushing process of m-sand fundamentally affects its shape characteristic, where the grading of m-sand and the proportion of micro fines may not be the same as fine aggregate. the percentage decrease in the compressive strength of 28 days immersed samples with respect to basic mix is estimated to be 2.08, 5.20, 8.86, 6.23 and 8.32% for 25, 50, 75 and 100% replacement of recycled coarse, respectively. notably, this current study was essentially carried out to devise the possibility of replacing the 100% sand by m-sand. using m-sand, the water demand may increase due to its high fines; however, the fines in m-sand play role to increase in paste volume, which will be useful in the development of scc with improved properties. 3.1.3. compressive strength of recycled coarse aggregate concrete and substitution of m-sand in fine aggregate the compressive test results of the recycled coarse aggregate replaced in natural coarse aggregate and replacing m-sand in fine aggregate are shown in table 2. from the obtained results on compressive strength, it can be noticed that with increasing replacement of recycled aggregate and m-sand, the effectual compressive strength is decreasing continuously in 7 and 28 days immersed samples. however, it is worth to mention that the researchers were widely used different waste materials other than the m-sand in fine aggregate. the percentage decrease in the compressive strength for 7 days immersed samples with respect to reference mix was found to be 8.93, 14.91, 16.42 and 28.35% corresponding to 25, 50, 75 and 100% replacement of recycled coarse aggregate and m-sand in fine aggregate. similarly, the percentage decrease in the effective compressive strength for 28 days immersed samples with respect to reference mix was estimated to be 9.37, 15.61, 19.78 and 30.21% corresponding to 25, 50, 75 and 100% replacement of recycled coarse aggregate and m-sand in fine aggregate. however, for the mixture having 100% of recycled aggregate and 100% of m-sand, the compressive strength value was estimated to be 29.77 mpa, which is almost equal to m 25 grade concrete. therefore, it is suitable and can possibly be recommended to utilize 100% recycled aggregate and m-sand in producing the scc with improved properties. 94 vol. 33/2022 analysis of concrete and cement epd s. no mix type average strength (mpa) m0 m25 m50 m75 m100 1 r0 3.65 3.61 3.56 3.44 3.4 2 r25 3.49 3.45 3.42 3.34 3.25 3 r50 3.38 3.36 3.33 3.27 3.15 4 r75 3.25 3.22 3.19 3.16 3.05 5 r100 3.19 3.16 3.13 3.05 2.94 table 3. estimated split tensile strength of the various concrete mixtures. mix m0 m25 m50 m75 m100 remarks astm c1202 charge passed (coulomb) chloride penetrability r0 1155.00 1197.00 1246.00 1274.00 1316.00 > 4000 high r25 1217.00 1259.00 1300.00 1318.31 1357.00 2000 − 4000 moderate r50 1273.00 1310.00 1350.12 1368.24 1404.00 1000 − 2000 low r75 1326.45 1370.00 1400.00 1419.54 1446.00 100 − 1000 very low r100 1384.00 1420.00 1449.41 1470.90 1491.00 < 100 negligible table 4. estimated rcpt value of scc mixtures after 56 days. 3.2. split tensile strength 3.2.1. split tensile strength of rca in nca concrete mix with 0% m-sand from test results of split tensile strength, it can clearly be noticed that the increasing replacement of recycled aggregate continuously leads to the decrement in the split tensile strength as listed in table 3. as per the literature reports, researchers typically observed a decrement in the strength when the % of recycled aggregate is increased. however, m. seethapathi et al [12] encountered a kind of reverse trend and reported that the split tensile strength were increased with increasing percentage of the recycled aggregate. but there were no specific reasons mentioned for the observed reverse trend in their studies. in this present study, the mixture r0m0 was taken as a reference mix for all the mixtures. the split tensile strength value of the 28 days immersed 100% recycled coarse aggregate was found to be 3.19 n/mm2. it was observed that the split tensile strength was noted to be decreasing with increasing recycled aggregate replacements in 28 days immersed samples. this could be ascribed to the low quality of recycled aggregate as compared to the natural coarse aggregate. the decrement % in the split tensile strength for 28 days immersed samples with respect to reference mix was estimated to be 4.58, 7.98, 12.30 and 14.42% for 25, 50, 75 and 100% replacement of recycled coarse, respectively. 3.2.2. split tensile strength of m-sand in fine aggregate with 0% recycled coarse aggregate the split tensile strength of m-sand replaced in fine aggregate samples is displayed in table 3. it can be noticed that the increasing replacing percentage of m-sand, the split tensile strength is continuously decreasing. the split tensile potency of 28 days immersed 100% m-sand in fine aggregate was estimated to be 2.81 n/mm2. it can clearly be noticed that the split tensile is decreasing with increasing rca replacement ratios in 28 days immersed samples. this could be due to the relative improved quality of msand that varies with river sand. the percentage decrease in split tensile for 28 days immersed materials with respect to the reference mixture r0m0 is found to be 1.10, 2.52, 6.10 and 12.30 corresponding to 25, 50, 75 and 100% replacement of m-sand in fine aggregate. 3.2.3. split tensile strength of the recycled coarse concrete and replacement of m-sand in nca and fine aggregate the split tensile strength results of the recycled coarse aggregate replaced in the natural coarse and m-sand replaced fine aggregate are presented table 3. it can be noticed that the increasing replacing % of recycled coarse aggregate and fine aggregate by msand continuously decreases the split tensile strength in 28 days immersed samples. notably, no literature reports are available studying the replacement of rca in nca and m-sand in fine aggregate. the percentage decrease in split tensile for 28 days immersed samples with respect to the reference mixture r0m0 is found to be 5.79, 9.60, 15.50 and 31.76% corresponding to 25, 50, 75 and 100% replacement of recycled coarse aggregate and m-sand in fine aggregate samples. however, it is noteworthy that the mixture having 100% recycled aggregate and 100% m-sand, the split tensile value was estimated to be 2.77 mpa. therefore, it can possibly be recommended to utilize the 50% recycled aggregate and m-sand in producing the scc towards having the improved split tensile 95 s. chundupalle, n. seeri, s. r. veera, v. r. nelluri acta polytechnica ctu proceedings mix m0 m25 m50 m75 m100before after before after before after before after before after r0 29.77 26.40 28.88 25.46 26.22 23.06 25.33 22.09 24.00 20.90 r25 27.74 24.17 26.94 23.46 24.94 21.65 24.00 20.80 22.67 19.62 r50 25.77 22.38 24.94 21.60 23.21 20.04 22.54 19.41 21.34 18.37 r75 24.00 20.75 23.21 20.04 21.67 18.70 20.93 18.02 19.94 17.10 r100 22.27 19.18 21.40 18.42 20.07 17.25 19.53 16.75 18.74 16.05 table 5. compressive strength of scc mixtures after 56 days of acid immersion. strength. 3.3. rapid chloride permeability test (rcpt) table 4 displays the rapid chloride (cl) permeability characteristics of scc mixtures after 56 days cured samples. from the table 4, it can be noticed that the increasing percentage of recycled aggregate leads to the increment in the rcpt values of scc mixtures. the permeability value of scc mixture without any recycled coarse aggregate replacement was found to be 1155 and with 100% replacement, the value was increased to 1384. the decrease in permeability values could be attributed to the insufficient properties of the recycled coarse s as compared to the natural coarse aggregate. it could also be observed that with increasing replacements of m-sand in fine aggregate, the permeability is also increasing in the samples. accordingly, the permeability value of the mixture with 100% replacements of fine aggregate with msand was estimated to be 1316. it is also observed that permeability is found to be increasing when the percentage of recycled coarse and m-sand is replaced at a time in natural coarse aggregate and fine aggregate. the permeability value with 100% replacement of recycled coarse aggregate and m-sand was found to be 1491. 4. conclusions from this current study and the obtained results, the following conclusions could be arrived; • from the obtained results of the present investigation, it can be recommended that the recycled coarse aggregate and m-sand can be effectively used in developing scc with relatively improved properties. • the compressive strength of scc is decreasing with increasing % replacements of recycled coarse aggregate in natural coarse aggregate and m-sand in fine aggregate. the compressive strength of the mixture having 100% recycled aggregate and 100% m-sand was estimated to be 29.77 mpa, which is almost equal to m 25 grade concrete’s strength • the split tensile strength of scc is decreasing with increasing percentage replacement of recycled coarse aggregate in natural coarse aggregate and m-sand in fine aggregate. the split tensile strength of the mixture having 100% recycled aggregate and 100% m-sand was estimated to be 2.77 mpa. • permeability values are found to be increasing as the replacement of recycled coarse aggregate and m-sand in natural coarse aggregate and fine aggregate are increasing. • when immersed in hcl, the compressive properties of scc were decreasing when the replacements of recycled coarse aggregate and m-sand in natural coarse aggregate and fine aggregate are increasing. • from the strength properties observed in this study, the scc can be effectively produced with recycled coarse aggregate replacement as well as m-sand replacement. references [1] j. m. v. gómez-soberón. porosity of recycled concrete with substitution of recycled concrete aggregate. cement and concrete research 32(8):1301-11, 2002. https://doi.org/10.1016/s0008-8846(02)00795-0. [2] l. evangelista, j. de brito. mechanical behaviour of concrete made with fine recycled concrete aggregates. cement and concrete composites 29(5):397-401, 2007. https: //doi.org/10.1016/j.cemconcomp.2006.12.004. [3] h. okamura, m. ouchi. self-compacting concrete. journal of advanced concrete technology 1(1):5-15, 2003. https://doi.org/10.3151/jact.1.5. [4] s. c. kou, c. s. poon. properties of self-compacting concrete prepared with coarse and fine recycled concrete aggregates. cement and concrete composites 31(9):622-7, 2009. https: //doi.org/10.1016/j.cemconcomp.2009.06.005. [5] z. j. grdic, g. a. toplicic-curcic, i. m. despotovic, et al. properties of self-compacting concrete prepared with coarse recycled concrete aggregate. construction and building materials 24(7):1129-33, 2010. https: //doi.org/10.1016/j.conbuildmat.2009.12.029. [6] p. nanthagopalan, m. santhanam. fresh and hardened properties of self-compacting concrete produced with manufactured sand. cement and concrete composites 33(3):353-8, 2011. https: //doi.org/10.1016/j.cemconcomp.2010.11.005. [7] k. c. panda, p. k. bal. properties of self compacting concrete using recycled coarse 96 https://doi.org/10.1016/s0008-8846(02)00795-0 https://doi.org/10.1016/j.cemconcomp.2006.12.004 https://doi.org/10.3151/jact.1.5 https://doi.org/10.1016/j.cemconcomp.2009.06.005 https://doi.org/10.1016/j.conbuildmat.2009.12.029 https://doi.org/10.1016/j.cemconcomp.2010.11.005 vol. 33/2022 analysis of concrete and cement epd aggregate. procedia engineering 51:159-64, 2013. https://doi.org/10.1016/j.proeng.2013.01.023. [8] s. santos, p. da silva, j. de brito. mechanical performance evaluation of self-compacting concrete with fine and coarse recycled aggregates from the precast industry. materials 10(8), 2017. https://doi.org/10.3390/ma10080904. [9] e. gueyisi, m. gesoglu, z. algin, et al. rheological and fresh properties of self compacting concrete containing coarse and fine recycled concrete aggregate, cement and concrete composites 113:622-630, 2016. [10] m. m. kamal, m. a. safan, z. a. etman, et al. behavior of recycled self-compacting concrete. erj engineering research journal 36(3):317-27, 2013. https://doi.org/10.21608/erjm.2013.67433 . [11] m. k. mounir, a. s. mohamed, a. e. zeinaband, et al. effects of polyprophelene fibers on development of fresh and hardened properties of recycled self compacting concrete, international journal of engineering and advanced technology 2(5):86-94, 2013. [12] a. tamilarasan, s. sankaran. experimental investigation of mechanical and durability properties of self compacting concrete with fly ash and msand, national conference on recent advances in coomunication, computation, electrical science and structures (ncraccess 2015), p. 11-16, 2015. [13] m. seethapathi, s. r. r. senthilkumar, k. chinnaraju. experimental study on high performance self compacting concrete using recycled aggregate, journal of theoretical and applied information technology 67(1):84-90, 2014. 97 https://doi.org/10.1016/j.proeng.2013.01.023 https://doi.org/10.3390/ma10080904 https://doi.org/10.21608/erjm.2013.67433 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0116 acta polytechnica ctu proceedings 38:116–123, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague integrated sustainability assessment using bim maria teresa henriques alves ferreiraa, ∗, antónio aguiar costab, josé dinis silvestrea a university of lisbon, instituto superior técnico, civil engineering, ceris, av. rovisco pais 1, 1049-001, lisbon, portugal b built colab, collaborative laboratory for the digital built environment, rua de álvares cabral 306, 4050-041 oporto, portugal ∗ corresponding author: teresaferreira@tecnico.ulisboa.pt abstract. the construction industry is responsible for 40 % of the energy consumption and 36 % of the co2 emissions, and buildings are responsible for a significant part of energy consumption in europe. thus, a growing concern regarding environmental impacts in the construction sector is in place. reducing these impacts and optimise the design process is a major priority, and technology needs to be integrated along with the design to allow for better buildings performance. building information modelling (bim) methodology is one of the technologies that is revolutionising how the supply chain delivers the construction projects, allowing for an overview of the whole life cycle, keeping track of the data along the process, and potentiating more advanced simulations and supported decisions. the tool proposed in this paper aims to integrate different types of sustainability analysis, namely streamlined life cycle assessment (lca), carbon footprint, life cycle cost (lcc) and level(s) framework with bim. this involves defining adequate product data templates (pdt) and a database structure for bim objects, including the necessary parameters to enable designers to do holistic and dynamic assessments from early design stages to a complete lca. also, considering the importance of using bim to visualise different scenarios, a graphical interface will be developed to show the key sustainability indicators and support decision-making for more sustainable buildings. the results achieved show that technology must be taken to meet climate most ambition targets and reduce the impact of construction. keywords: building information modelling, sustainability, technology, life cycle assessment, life cycle cost. 1. introduction in line with the most ambitious targets established in paris agreement in 2015, eu green deal has an ambition of reducing carbon emissions by 55 % or more until 2030 and for europe to be a climate-neutral continent by 2050 [1, 2]. construction and retrofit of buildings cause substantial environmental impacts [3] due to their significant consumption of energy (40 %) and materials and energy-related greenhouse gas emissions (36 %) [4]. the application of environmental (lca, including global warming potential impact category) and economic (or life cycle cost – lcc) life cycle assessment (lca) [5] to construction are methodologies that: are recognized and standardized at the european level [6, 7], are becoming more and more important, and which are being increasingly used in this sector by experts and researchers. the life cycle (lc) paradigm emerges, and construction materials and buildings now represent the sum of all impacts and costs in the respective lc (no longer seen as individual impact and cost). the construction industry plays a major role in the decarbonization process [3]. however, the lack of interoperability between eu targets for sustainability assessment and supply chain tools is a major issue in archive such important goals on climate issues. in figure 1 is shown a diagram of circular economy principles as a circle of interaction covering all lc stages of its analysis. applying circular economy principles (figure 1), current research aims to use sustainability indicators established in level(s) framework in a new construction case study. this is made by using a bim environment to complete both economic and environmental assessment analyses. circular economy principles and bim (figure 2) have a common ground in their full lc approach, and this is seen by the authors as a huge advantage that should be merged contributing to digitalization and circularity development. level(s) is a common framework for sustainable buildings assessment across europe, that has a whole lc approach. it considers measurement and improvement from design to end-of-life, covering both renovation and new construction [8]. level(s) framework provides a common methodol116 https://doi.org/10.14311/app.2022.38.0116 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 integrated sustainability assessment using bim figure 1. circular economy principles. ogy for assessing the sustainability of buildings based on six macro-objectives. it contributes therefore to achieving eu and member states policy goals in energy, material use and waste, water, and indoor air quality in an lc perspective [8]. intending to bring buildings into the circular economy, level(s) comprises a set of indicators, scenario tools, a data collection tool, checklists, and rating systems that allow professionals and project actors to measure buildings’ performance [9]. the aim of this research is to aggregate green ambitions with technology by using building information modelling (bim) software to perform building environmental impacts and help supply chains to make their decisions through the full lc of the building, covering from early design stages decisions to the end of life. this was done by developing an lca plugin for revit software to perform lca and lcc calculations, the plugin creates a set of parameters, each one representing an environmental or economic indicator, and does the calculation using the information and quantities from the bim model. 2. materials and methods in order to answer research questions and meet the defined goals, the plugin proposed integrates with bim different types of sustainability analysis in its functionalities, namely streamlined lca, carbon footprint, and lca in level(s) framework. economic analysis is also important and the parameter to measure lcc is also integrated into the plugin. to do so, it was necessary to develop an adequate product data template (pdt) for bim objects and construction elements, including the necessary parameters to enable designers to do holistic and dynamic assessments from early design stages to a complete lca at the end of the design process. the objective is to give the user a set of options to perform environmental and economic calculations, nowadays there are different ways to archive so, the aim is to give the user different functionalities that correspond to different ways to figure 2. bim – full life cycle approach. get the lca results for the user to choose the best one that fits its project objectives. to perform an lca based on level(s) framework analysis, the corresponding indicators were studied to find which ones can be read within a bim environment, in bim object or elements, and possibly converted into parameters of analysis in the scope of lca studies. from the six level(s) macro objectives, the ones necessary for an lca analysis were selected: objective 1 – greenhouse gas emissions along building life cycle, covering lc global warming potential; and objective 6, optimized lcc and value. from the analysis of these macro-objectives, the following indicator was considered to do level(s) lca: global warming potential (gwp) measuring the co2 emissions. to go further on the cradle to grave lca analysis as recommended by eu [10] on the standard en15804:2012+a2:2019 [11], the following indicators were also addressed: abiotic depletion potential for fossil fuels (adpe), abiotic depletion potential of materials (adpm), acidification potential (ap), eutrophication potential (ep), ozone depletion potential (odp), photochemical ozone creation potential (pocp), water (user) deprivation potential, deprivation-weighted water consumption (wdp). to reach macro-objective 6 optimized life cycle cost and value indicator, product and construction stages, use stage, and end of life stage, containing parameters for initial costs, annual costs, and periodic costs, were considered to measure lcc. 2.1. plugin functionalities mentioned indicators are read in the construction bim elements/objects and can be analyzed directly within a bim software. the bim-based environment chosen to apply the plugin was autodesk revit. plugin functionalities were divided into four main topics, following the areas of research; (i) streamlined lca; (ii) carbon footprint; (iii) level(s) lca; (iv) cost. 117 m. t. h. a. ferreira, a. aguiar costa, j. d. silvestre acta polytechnica ctu proceedings figure 3. plugin functionalities with indicators. in figure 3 is possible to find core indicators per category of analysis. those indicators will be converted to parameters to be read by revit shared parameters. 2.2. parametrization and product data template (pdt) the definition of an adequate pdt [12] structure for bim objects was necessary to correctly convert indicators to parameters to be read in a bim environment. table 1 shows the proposed pdt including the new shared parameters with the objective of allocating the data need to perform the lca calculations and analysis. 2.3. framework for bim lca lcc analysis tables 2–4 show lc stages for production, construction, use and end of life from a to d modules. the analytical calculation of the environmental impacts of the project on the production and construction phase (a1–a5 modules) is presented in equation (1) [13]: streamlined lca a1–a3︷ ︸︸ ︷ ei m cx = i∑ a=1 (qma × ei m a ) + complete lca︷ ︸︸ ︷ a4︷ ︸︸ ︷ j∑ b=1 (dvb × n × ei v b ) + a5︷ ︸︸ ︷ k∑ c=1 (qconc × ei con c ) + l × [a1 − a3], (1) ei m cx environmental impact of category x resulting from the manufacturing and construction phase (a1–a5 modules); i, j, k number of existing materials i, transportation j, and construction utilities k; qma quantity of material a; ei ma environmental impact (of category x) of material a; dvb distance from the supplier to the construction site, multiplied by the number of travels (n); ei vb environmental impact (of category x) of transportation b; qconc consumption of utility c throughout the construction (e.g., electricity [kwh], gas [mj], water [m3]); ei conc environmental impact (of category x) of utility c; l percentage of wasted materials during the construction phase. the analytic calculation of the economic impacts of the project, on the production and construction phase 118 vol. 38/2022 integrated sustainability assessment using bim indicator/ parameter name streamlined lca complete lca level(s) lca carbon footprint level(s) lcc bim objects/ elements project details units adpe x x x x mj adpm x x x x kg sb eq pe-re x x x mj pe-nre x x x mj ap x x x x kg so2 eq ep x x x x kg po3−4 eq odp x x x x kg r-11 eq pocp x x x x kg c2h4 wdp x x m 3 world eq deprived gwp x x x x x kg co2 eq lcc-ic x x €/m2/yr lcc-ac x x €/m2/yr lcc-pc x x €/m2/yr lcc-eol x x €/m2/yr durability x x yr density x x ρ = m/v building’s service life x x yr transportation type x x transportation distance x x km utilities used x x x waste generated x x x table 1. proposed pdt. raw material supply transport manufacturing construction installation production a1 a2 a3 construction a4 a5 table 2. lc stages a modules/production & construction stages. use maintenance repair replacement refurbishment operational operationalenergy use water use use stage b1 b2 b3 b4 b5 b6 b7 table 3. lc stages b modules/use stage. deconstruction, demolition transport waste processing disposal reuse, recycling, or energy recovery potentials end of life stage c1 c2 c3 c4 d table 4. lc stages c/d modules / end of life. 119 m. t. h. a. ferreira, a. aguiar costa, j. d. silvestre acta polytechnica ctu proceedings (a1–a5 modules) is presented in equation (2): streamlined lcc a1–a3︷ ︸︸ ︷ c m c = i∑ a=1 (qma × ac m a ) + complete lcc︷ ︸︸ ︷ a5︷ ︸︸ ︷ j∑ b=1 (qmb × c c2b b ) + k∑ c=1 (qconc × c con c ) + l × [a1 − a3], (2) c m c costs resulting from the manufacturing and construction phase (a1–a5 modules); i, j, k number of existing materials i, j and construction utilities k; qma , qmb quantity of materials a or b used in the construction; ac ma the acquisition cost of material a; c c2bb cost to build/assemble construction elements (e.g., cost to apply 1 m 2 of mortar); qconc consumption of utility c throughout the construction; c conc cost of utility c; l percentage of wasted materials during the construction phase. 3. results in this article was developed an innovative approach for a level(s) lca methodology analysis and its integration into a bim plugin. in this way, designers are enabled to perform a holistic and dynamic assessment from early design stages to a complete lca, and rapidly visualize different scenarios, using core sustainability indicators. the innovative nature of the approach of the current plugin is that level(s) lca framework is integrated into bim and gives the users the possibility to choose which type of analysis they will want to perform, including streamlined lca, carbon footprint, level(s) lca and cost (lcc) (figure 4). also, considering the importance of using bim to visualize different scenarios, a graphical interface was developed to show the key sustainability indicators and support decision-making for more sustainable buildings. 3.1. buildings assessment information an example of environmental information read in revit regarding exterior double brick walls with insulation is illustrated in figure 5. it is possible to see the shared parameters for the lca analysis inserted in revit. figure 7 explains the workflow of the lca plugin, per analysis, and shows the connections, once one analysis is finalized, to go further/ deeper in the lca analysis. using the case study of a new family swelling, the authors ran the plugin for exterior walls. according to the workflow in figure 7, the first step is to choose the stage of the project development, either conceptual design (level 1) or detail design (level 2) stages. if level 2 is chosen, then the user is asked which functionality of the plugin he would like to perform (figure 4). if he chooses, for example, level(s) lca, the shared parameters corresponding to the environmental indicator of analysis, table 1, are loaded into the bim model and the data from the database is loaded to the project to be assembled to the correspondent elements, for example, a wall, figure 5. the environmental data is stored in the properties of the bim family and read also in the plugin. the next step is to perform the corresponding calculations, which are done automatically by the plugin following analytic models presented in section 2.3; for this example equation (1) is being used. results can be exported both to excel and 3d visualization with a colour scheme, as presented in figure 6 for the exterior walls sample. 4. discussion the potential of disruptive technologies, such as building information modelling (bim), changes how the assessment of its sustainability is conducted [14]. lc existing tools/platforms for bim available in the market that integrate sustainability analysis for building design were already analyzed by the authors, i.e. tally and one click lca. from that analysis, several gaps, lack of interoperability between tools, limited parameters of analysis, limited databases, and lack of bim libraries were found. regarding the digitalization of construction, the development of the proposed methodology and tool is fundamental to supporting the performance-driven architectural design of buildings [15]. following the results presented, the authors believe that a user-friendly tool to make simulations of different sustainability analyses helps designers and supply chains to make choices with lower environmental impact. the time and effort of level(s) lca are reduced, and a graphic user interface would provide the decision-making support for the design of sustainable buildings. for future research, the idea is to build a plugin that is flexible enough to address new indicators and analysis in its functionalities and also remove the 120 vol. 38/2022 integrated sustainability assessment using bim figure 4. plugin functionalities options. figure 5. environmental impacts incorporated in shared parameters of a wall in revit. figure 6. level(gwp) analysis complete 3d results for a wall. 121 m. t. h. a. ferreira, a. aguiar costa, j. d. silvestre acta polytechnica ctu proceedings figure 7. workflow scheme of the lca plugin. ones that with the following developments will be made redundant. it is also possible to expand the proposed methodology to assess other methods such as leed, bream, wells, or others but some of them have methods of analysis not quantitative and that is a limitation to where this plugin can automatically grow and specified engineering consultants is required to fulfill all the requirements of specified schemes. 5. conclusion the application of bim can simplify level(s) implementation by systematizing the approach and accelerating access to data. the integration of different types of sustainability analysis, namely streamlined lca, carbon footprint, lcc, lca/ level(s) framework, and lcc with bim enable designers and the construction industry to easily integrate and see the environmental impacts of their buildings, and possible study better alternatives and actively collaborate to reduce the impact of the construction industry on the important climate change issues. acknowledgements the authors acknowledge the civil engineering research and innovation for sustainability (ceris) research unit, funded by fundação para a ciência e a tecnologia (fct) and the scholarships attributed to the first author with the references ui/bd/153398/2022. the authors would also like to thank the support through the circular ecobim project, funded by eea grants within the environment programme. finally, the authors would like to demonstrate their gratitude to the design office “atelier dos remédios” and prof. dr. francisco t. bastos for providing a case study. references [1] ec. communication from the european commission: a renovation wave for europe – greening our buildings, creating jobs, improving lives, 2020. com(2020)662. brussels, belgium. [2] ec. renovation wave, 2021. [2021-03-11]. https://ec.europa.eu/energy/topics/energyefficiency/energy-efficientbuildings/renovation-wave_en [3] ec. communication from the european commission: the european green deal, 2019. com (2019) 640, brussels, belgium. [4] ec. energy efficiency – revision of the energy performance of buildings directive, 2010. [2020-03-11]. https://ec.europa.eu/info/law/betterregulation/have-your-say/initiatives/12910revision-of-the-energy-performance-ofbuildings-directive-2010-31-eu [5] r. santos, a. a. costa, j. d. silvestre, et al. bim-based life cycle assessment and life cycle costing of an office building in western europe. building and environment 169:106568, 2020. https://doi.org/10.1016/j.buildenv.2019.106568 122 https://ec.europa.eu/energy/topics/energy-efficiency/energy-efficient-buildings/renovation-wave_en https://ec.europa.eu/energy/topics/energy-efficiency/energy-efficient-buildings/renovation-wave_en https://ec.europa.eu/energy/topics/energy-efficiency/energy-efficient-buildings/renovation-wave_en https://ec.europa.eu/info/law/better-regulation/have-your-say/initiatives/12910-revision-of-the-energy-performance-of-buildings-directive-2010-31-eu https://ec.europa.eu/info/law/better-regulation/have-your-say/initiatives/12910-revision-of-the-energy-performance-of-buildings-directive-2010-31-eu https://ec.europa.eu/info/law/better-regulation/have-your-say/initiatives/12910-revision-of-the-energy-performance-of-buildings-directive-2010-31-eu https://ec.europa.eu/info/law/better-regulation/have-your-say/initiatives/12910-revision-of-the-energy-performance-of-buildings-directive-2010-31-eu https://doi.org/10.1016/j.buildenv.2019.106568 vol. 38/2022 integrated sustainability assessment using bim [6] cen. sustainability of construction works – assessment of buildings part 2: framework for the assessment of environmental performance. en 15643-2., 2011. comité européen de normalisation, brussels, belgium. [7] cen. sustainability of construction works – assessment of buildings part 4: framework for the assessment of economic performance. en 15643-4, 2012. comité européen de normalisation, brussels, belgium. [8] ec. level(s), putting whole life carbon into practice. publications office of the european union, 2021. https://doi.org/10.2779/79139 [9] ec. level(s), putting circularity into practice. publications office of the european union, 2021. https://doi.org/10.2779/19010 [10] n. dodd, s. donatello, m. cordella. level(s) indicator 1.2: life cycle global warming potential user manual: overview, instructions and guidance (publication version 1.0), 2020. [11] en15804:2012+a2:2019, sustainability of construction works – environmental product declarations – core rules for the product category of construction products. [12] m. signorini, s. frigeni. development of a bim-based product data template for sustainability. master’s thesis, politecnico di milano, arc i – scuola di architettura urbanistica ingegneria delle costruzioni, 2018. http://hdl.handle.net/10589/143041 [13] r. santos, a. a. costa, j. d. silvestre, l. pyl. integration of lca and lcc analysis within a bim-based environment. automation in construction 103:127–149, 2019. https://doi.org/10.1016/j.autcon.2019.02.011 [14] r. santos, a. a. costa, j. d. silvestre, l. pyl. informetric analysis and review of literature on the role of bim in sustainable construction. automation in construction 103:221–234, 2019. https://doi.org/10.1016/j.autcon.2019.02.022 [15] x. shi, w. yang. performance-driven architectural design and optimization technique from a perspective of architects. automation in construction 32:125–135, 2013. https://doi.org/10.1016/j.autcon.2013.01.015 123 https://doi.org/10.2779/79139 https://doi.org/10.2779/19010 http://hdl.handle.net/10589/143041 https://doi.org/10.1016/j.autcon.2019.02.011 https://doi.org/10.1016/j.autcon.2019.02.022 https://doi.org/10.1016/j.autcon.2013.01.015 acta polytechnica ctu proceedings 38:116–123, 2022 1 introduction 2 materials and methods 2.1 plugin functionalities 2.2 parametrization and product data template (pdt) 2.3 framework for bim lca lcc analysis 3 results 3.1 buildings assessment information 4 discussion 5 conclusion acknowledgements references https://doi.org/10.14311/app.2022.33.0168 acta polytechnica ctu proceedings 33:168–174, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague fire spalling sensitivity of concrete made with recycled concrete aggregates (rca) bruno fernandesa, ∗, hélène carréa, jean-christophe mindeguiab, céline perlota, christian la borderiea a université de pau et des pays de l’adour, e2s uppa, siame, allée du parc montaury, 64600 anglet, france b université de bordeaux, laboratoire l2m, cours de la libération, 33 405 talence, france ∗ corresponding author: bruno.fernandes@univ-pau.fr abstract. the fire spalling of concrete is a complex phenomenon, which can affect the integrity of the structures during a fire. this thermal instability is associated with a complex coupled chemo-thermohydro-mechanical mechanism and it can be influenced by many factors, related to material (e.g. permeability, porosity and water content), geometry (e.g. shape and size) and environmental parameters (e.g. mechanical load and heating rate). concrete made with recycled concrete aggregates presents higher porosity, higher water content and different interfaces between aggregates and mortar. all these aspects can lead to a different behaviour under fire exposure, including the spalling risk of these sustainable concretes. the main objective of this paper is to analyse the influence of the use of recycled concrete aggregates on the spalling risk of concrete. in this paper, concrete prisms with different replacement rates of recycled coarse aggregates (0 up to 100%) were exposed to a standard fire curve (iso 834-1) with a constant uniaxial compression load. after heating, samples surfaces were evaluated by means of digital photogrammetry. results showed that concrete with rca is sensitive to explosive spalling. all replacement rates presented higher degree of spalling than concrete made with natural aggregates. keywords: fire, recycled concrete aggregate, spalling. 1. introduction in the last decades, the use of recycled concrete aggregates (rca) has gained popularity due to their potential to make sustainable concrete. as for conventional concretes, the fire behaviour of these concretes, including their spalling sensibility, must be well known to guaranty the safety of the structure in case of a fire. moreover, fire behaviour is fundamental to enhance the resilience of the buildings for these types of extreme events. fire spalling is the detachment of concrete fragments from structures exposed to fire. this phenomenon may affect the integrity of structures, reducing thermal insulation, exposing the steel rebar to fire and reducing the load-bearing capacity [1, 2]. despite the numerous experimental and numerical studies, there is still a lot of uncertainties about the physical origin of the phenomenon [2]. in general, spalling is associated with a coupled thermo-hygrochemo-mechanical mechanism. in general, spalling occurs in different forms, such as aggregate spalling, surface spalling, explosive spalling and corner spalling [3]. previous studies show that many parameters can influence the occurrence of spalling, and it can be divided into material (e.g. permeability, porosity and water content), geometry (e.g. shape and size) and environmental parameters (e.g. heating rate and mechanical load) [3, 4]. the susceptibility to spalling of concrete made with recycled aggregates (rca) is barely known. since rca aggregates are more porous, the use of this kind of aggregates can contribute to vapour migration, reducing spalling risk [5]. at the same time, these concretes present higher water content and different phase interfaces, which also may affect the spalling risk. previous works verified occurrence of this phenomenon. robert et al. [6] observed superficial and localized spalling in full-scale slabs made with rca. authors attributed this behaviour to the high water content of the mix. pliya et al. [5] observed spalling on some cylinder samples exposed to a rapid heating rate. it is noteworthy that in this case, samples are classified as high strength concrete, which is more prone to spalling due to its denser microstructure. the aim of this study is then to evaluate the influence of recycled concrete aggregates (rca) on the spalling behaviour. to this, uniaxial spalling tests were carried out on of ordinary concrete samples containing different rca (coarse aggregates) replacement rates (0%, 10%, 20%, 40%, 70% and 100%). during fire, spalling events are followed and after fire, the residual sample geometry was evaluated by digital photogrammetry, in terms of volume and spalling 168 https://doi.org/10.14311/app.2022.33.0168 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 fire spalling sensitivity of rca concrete depth. 2. experimental program 2.1. concrete mixes cement cem ii/a-l 42.5 r ce was used. to achieve workability, limestone filler betocarb hpsc and sika viscocrete tempo-483 superplasticizer were also used in the mixes. sand used in this study had a fineness modulus of 3.10, density of 2650 kg/m3 and water absorption of 0.35 %. two types of coarse aggregates were used: natural aggregates (na) and recycled concrete aggregates (rca). for each type of aggregates, two fractions were used, 4/10 and 10/20. natural aggregates were made from diorite and have a density of 2820 kg/m3 (4/10) and 2840 kg/m3 (10/20), and a water absorption of 0.92 % (4/10) and 0.81 % (10/20). recycled concrete aggregates were crushed and screened by a recycling company, and have a density of 2570 kg/m3 (4/10) and 2590 kg/m3 (10/20), and a water absorption of 5.60 % (4/10) and 4.52 % (10/20). as expected, rca present higher water absorption values, mainly due to the presence of adhered "old" mortar. rca aggregates also present some impurities, such as wood, plastic, brick and bituminous materials. six different ordinary concrete mixes were prepared and mix proportions are given in table 1. for concretes made with rca, na were directly replaced by rca in volume, without any changes in the other components. samples were designed to meet nf en 206/cn:2014 [7] durability requirements for xd3 exposure class. to compensate the higher water absorption of rca and aiming a practical method that can be reproduced in a real concrete plant, all aggregates (sand, na and rca) were pre-wetted for one hour using a soaker hose and then covered with plastic sheet for three hours. all mixes have the same water/cement ratio (0.5) and dosage of superplasticizer (0.9 % c). to avoid errors when determining water content in aggregates, the amount of water in the batch is adjusted to obtain approximatively the same slump (180 − 190 mm). slump, mechanical and physical properties for all studied mixes are reported in table 1. for each mix, mechanical properties at 28 days and at fire test age were measured on 11 × 22 cm and 16 × 32 cm cylindrical specimens. for spalling tests, 20 × 20 × 10 cm (height x length x thickness) prisms were casted. cylindrical samples used to determine 28 days properties were kept submerged in water at 20 řc until test day. samples for fire tests were first kept submerged for seven days and then placed into sealed plastic bags until fire test age (between 88 and 125 days). this curing induces high water content, thus inducing high probability of spalling for all the mixes. at test age, water content (wc), was also measured. as expected, higher the replacement rate, higher the water content in the samples, due to higher porosity of rca. from na to rca-100, a variation of 2.11% is observed. the influence of the water content on the different spalling behaviour between concretes made with na or rca are discussed in section 3. 2.2. spalling evaluation spalling was evaluated by testing concrete prisms subjected to a standard fire (iso 834-1) and uniaxial loading. three samples per concrete mix were tested (with exception of rca-10). before the tests, the sides of the prisms were sealed with aluminium foil tape (figure 1a), which was glued using high temperature glue. this was done to avoid moisture loss through the sides of the concrete blocks and provide unidirectional transport condition of heat and moisture. an intermediate scale furnace, powered by one propane gas burner, was used for the heating process (figure 1b). gas flow was controlled during the test in order to follow the iso 834-1 fire curve. the furnace has an opening of 20 cm x 20 cm and three type k thermocouples were used to measure the heating curve. these thermocouples were placed in the front of the furnace and one centimetre from the heated face of the sample, at three different heights: 4, 10 and 16 cm from the bottom of the furnace opening. more detail can be found in previous works [8, 9]. a hydraulic press was used to apply a uniaxial loading of 5 mpa, in order to be representative of a conventional structure (wall or column for example). the tested specimen was placed between two concrete blocks. the sides of the tested specimen and the support blocks were insulated with rock wool. figures 1c and 1d presents the full test setup. first, the load was applied and after stabilization, the heating process started. the iso 834-1 fire curve was applied for 30 minutes and the three thermocouples allowed to follow the temperature of air close to the sample. during tests, each spalling event has been registered, in terms of occurrence time and intensity (small, medium and big noise). after 30 minutes, heating was turned off and after 5 minutes, the sample was discharged. after heating, post-fire analysis was conducted on all samples with a digital photogrammetry technique. this method allows to construct 3d models from digital photos [10, 11]. the method follows an approach that involves image acquisition and processing, 3d reconstruction, mesh generation and post processing [10]. a digital camera (panasonic dmc-gx80) was used. samples were placed in a rotating table, in order to successfully orientate 3d reconstruction. camera was fixed on a tripod and between each shot, the rotating table was turned around 10ř. a total of 36 images were obtained for each sample. the image processing and 3d reconstruction were done with the meshroom [12] software and later imported 169 b. fernandes, h. carré, j.c. mindeguia et al. acta polytechnica ctu proceedings material / property mixna rca-10 rca-20 rca-40 rca-70 rca-100 cement/filler/sand (kg/m3) 350 / 60 / 804.3 na 4/10 (kg/m3) 331.7 298.5 265.3 199 99.5 na 10/20 (kg/m3) 711.1 640 568.9 426.7 213.33 rca 4/10 (kg/m3) 30.2 60.5 120.9 211.58 302.3 rca 10/20 (kg/m3) 64.9 129.7 259.4 453.96 648.5 water (kg/m3) 175 slump (mm) 194 184 174 176 158 210 fc28 (mpa) 44.8 (±0.2) 39.7 (±0.1) 37.4 (±3.5) 39.7 (±3.0) 34.5 (±2.7) 34.6 (±1.1) j (days) 94 116 88 88 88 125 fc,j (mpa) 46.6 (±1.4) 44.5 (±4.6) 50.8 (±0.5) 49.6 (±0.4) 45.9 (±0.3) 39.1 (±0.3) ec,j (gpa) 34.5 36.8 34.9 39 30.1 30.5 wcj (%) 4.94 (±0.4) 5.02 (±0.5) 5.10 (±0.2) 5.36 (±0.4) 6.41 (±0.5) 7.05 (±0.3) table 1. mix design and properties. figure 1. furnace details and test set-up. 170 vol. 33/2022 fire spalling sensitivity of rca concrete figure 2. developed fire curves. figure 3. first spalling event. in cloudcompare [13], where the model was cleaned and scaled. after this processing, a binary mesh was generated and used in a cast3m [14] algorithm, which post-processed the model to obtain the spalling volume and depths. the error from the methodology applied here is 4.2% (previously determined). 3. results 3.1. fire curves and spalling events fire curves developed during fire tests are presented in figure 2. in grey, the average of the three thermocouples is presented, together with the first registered spalling event. in general, curves follow properly the iso 834-1 fire curve and the perturbations start after the first spalling event. the oven temperature at the time of the first spalling is usually around 550 ◦c to 650 ◦c. the time of first spalling is also presented in figure 3, which presents the time of this event in relation to the replacement rate. except for rca-10 (indicated in figure 2), all samples first spalled between 5 and 10 minutes. the perturbations with rca-10 are related to the fact that the curve in this tests lasted figure 4. spalling events. figure 5. total number of spalling events. more to reach the temperature that may trigger the spalling. besides the first event, each spalling was recorded and categorized in three different qualitative "sound intensity": small, medium and big spalling. in order to have a visualization of these different events along time, figure 4 is proposed. in this figure, results from three representative samples were plotted. in general, all samples started with small spalling events (popup, bottom of the figure 4), then sound intensity increases with the fire duration. the last events usually had big intensity and happened around 10 − 15 minutes. figure 5 presents the total amount of spalling events for each of the tests in relation to the replacement rate. from the figure 5, we can see it a lot of discrepancies in the spalling events and no direct link with the quantity of recycled aggregates can be observed. 3.2. inspection after fire spalling test different analyses were done after the fire spalling tests. first, a visual inspection was done. examples 171 b. fernandes, h. carré, j.c. mindeguia et al. acta polytechnica ctu proceedings figure 6. samples after fire: photos and depth colour map. figure 7. max spalling depth. of samples exposed surface are presented in figure 6. visually it is already possible to observe an influence of the type of aggregates on the spalling severity: na samples present less spalling than mixes made with rca. the impurities in tested samples are also observed. especially in rca 100, the presence of bituminous particles is clearly seen, that burned and melted during the fire test. its noteworthy that the presence of this impurities reduces the strength of concrete made with rca at high temperatures, as observed in laneyrie et al. [15]. figure 6 also presents the colour maps generated in the photogrammetry process. max spalling depth and spalled volume were measured with photogrammetry for all the replacement rates (figures 7 and 8). the first aspect is that concrete made with na presents lower values of spalling depth and volume than any rca replacement rate. these graphs confirms some of the tendencies observed in the visual inspection, i.e. concrete made with na presents lower figure 8. spalled volume. values of spalling volume and depths. another observation is that the evolution of spalling volume and depth with the replacement rate is quite similar. in the first branch, from na to concrete made with 20% of rca, an almost linear increase is observed. in the case of the spalled volume, this increase also appears from rca-20 to rca-40, but the slope of increase is lower. from 40% to 100%, the average values keep quite constant, with no major changes, independently of the quantity of rca. a higher variability of results is also noteworthy for concretes made with rca. first, the use of rca results in a higher water content, as seen in table 1. usually, higher the water content (or the moisture content), higher is the spalling degree [16]. it seems then logical that the rca-100, which had the higher water content at the test age (7.05 %), experimented the more severe spalling (volume and depth). however, the influence 172 vol. 33/2022 fire spalling sensitivity of rca concrete of water content for the other mixes is not that clear. indeed, when comparing na and rca-20, the increase of spalling volume (5.51 %) is much higher than the increase in water content (0.16 %). moreover, it can be seen that spalling volume and depth are quite constant from rca-40 to rca-100, whereas the water content increases significantly (1.5 %). these factors indicate that the water content may influence the spalling of concrete made with rca, but is not the only parameter affecting the spalling sensibility. another influencing parameter for spalling is the compressive strength. usually, the higher the strength, the higher will be the spalling probability [16, 17], due to the lower permeability [16] of the mix. this is especially verified with hpc, where the spalling is a real concern [18]. yet, the influence of compressive strength on spalling is not clearly observed in our study. indeed, the mix with the lower compressive strength at test age (rca100, 39.1 mpa) presents the higher spalled volume. this shows that, in this experimental campaign, the strength was not the main factor affecting spalling sensitivity of concrete made with rca. lastly, internal thermal cracking also may have a contribution to the spalling events. usually, cracks increase the matrix permeability, allowing pressure release and moisture transport, and then a reduction of the spalling risk [1, 19]. these cracks are mainly due to thermal mismatch between aggregates (that expand) and cement matrix (that shrinks after 100 ◦c). the cracking of concrete made with rca under fire exposure is still not clear. on one hand, increasing the replacement rate of rca involves, for a fixed volume of new paste, a higher proportion of old mortar / new paste interfaces, and consequently a lower proportion of new paste / aggregates interfaces. this can probably reduce the density of thermal cracking in concretes made with rca, and then potentially increase the spalling sensibility. on the other hand, the use of rca also may lead to an excess of water (regarding only the one necessary for cement hydration), especially surrounding the aggregates themselves [5]. this extra water may reduce the bond with the new paste, inducing cracks and a local increase of the permeability [5]. this point will be further analysed by means of microscopic observations of heated samples, permeability measurement and thermal expansion measurements. 4. conclusions this paper presented an experimental investigation into the spalling sensitivity of concrete made with rca. based on the test results, the following conclusions are drawn: • visual inspection showed that concrete made with rca presents a spalling more significant than concrete made with na. • the time of the first spalling event can be a good indicator of the spalling occurrence. usually, this time is associated with the temperature reached in the material. in this study, no clear influence of the rca on the time of first spalling was observed, with a high variability of results. • the evolution of the spalling depth and spalling volume with the rca replacement rate both present a peculiar behaviour. first, an increase is observed (from na to rca-20), then values keep almost constant up to rca-100. both indicators showed that concretes made with rca presents a higher spalling sensibility. • different factors may influence the risk of spalling of concrete made with rca: water content, strength and thermal cracking. in this experimental campaign, compressive strength does not affect the spalling behaviour of concrete with rca, since concrete mixes with relatively low strength presented a high spalling sensibility. water content seems to affect the spalling sensibility, since mixes with high water content presented the higher spalling depth and volume. however this is not the main factor, since samples made with rca and without rca, and with similar water content, presented big variations of spalling depth and volume. thermal cracking should be further investigated, along with permeability measurements, microscopic observations and thermal expansion. • lastly, it is important to note that all these experimental conclusions are drawn based on a specific thermal and mechanical state, and for a given geometry of samples. to be more representative of real structures, full-scale fire tests on structural concrete elements made with rca are planned, in order to analyze (i) the spalling sensibility of these concretes in a real situation and (ii) the influence of this spalling extent on the fire resistance of the structure. acknowledgements the authors would like to thank région nouvelleaquitaine for funding support (project recyfeu). they also would like to thank the support of groupe cassous (guyenne environnement and aqio) and groupe garandeau. references [1] j.-c. mindeguia, p. pimienta, a. noumowé, et al. temperature, pore pressure and mass variation of concrete subjected to high temperature experimental and numerical discussion on spalling risk. cement and concrete research 40(3):477-87, 2010. https: //doi.org/10.1016/j.cemconres.2009.10.011. [2] j.-c. mindeguia, h. carré, p. pimienta, et al. experimental discussion on the mechanisms behind the fire spalling of concrete. fire and materials 39(7):61935, 2015. https://doi.org/10.1002/fam.2254. 173 https://doi.org/10.1016/j.cemconres.2009.10.011 https://doi.org/10.1002/fam.2254 b. fernandes, h. carré, j.c. mindeguia et al. acta polytechnica ctu proceedings [3] g. a. khoury. effect of fire on concrete and concrete structures. progress in structural engineering and materials 2(4):429-47, 2000. https://doi.org/10.1002/pse.51. [4] féderation internationale du béton. fire design of concrete structures structural behaviour and assessment (lausanne, switerland), fib bulletin no. 46, 2008. [5] p. pliya, d. cree, h. hajiloo, et al. high-strength concrete containing recycled coarse aggregate subjected to elevated temperatures. fire technology 55(5):1477-94, 2019. https://doi.org/10.1007/s10694-019-00820-0. 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[17] p. pimienta, n. taillefer, p. pimienta, et al. spalling of concrete: a synthesis of experimental tests on slabs. matec web of conferences 6, 2013. https://doi.org/10.1051/matecconf/20130601008. [18] p. pimienta, r. j. mcnamee, j.-c. mindeguia. hysical properties and behaviour of high-performance concrete at high temperature (springer international publishing), 2019. [19] e. w. h. klingsch. explosive spalling of concrete in fire phd thesis (eth zurich), 2014. 174 https://doi.org/10.1002/pse.51 https://doi.org/10.1007/s10694-019-00820-0 https://doi.org/10.1002/suco.201800196 https://doi.org/10.1016/j.jterra.2017.06.001 https://github.com/alicevision/meshroom http://www.cloudcompare.org/ http://www-cast3m.cea.fr/ https://doi.org/10.1016/j.conbuildmat.2016.02.056 https://doi.org/10.1016/j.firesaf.2006.10.006 https://doi.org/10.1051/matecconf/20130601008 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.35.0019 acta polytechnica ctu proceedings 35:19–22, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague construction 4.0 in the concept of the railway pavel krupík czech technical university in prague, faculty of civil engineering, department of construction management and economics, thákurova 2077/7, 166 29 prague 6, czech republic correspondence: pavel.krupik@fsv.cvut.cz abstract. the digital and technological innovations that have been developing massively in recent years are often referred to as the 4th industrial revolution or industry 4.0. the construction industry seems to be somewhat on the sidelines. the big topic is bim (building information modelling), but this is only one component of the contemplated construction 4.0. other topics, if not completely neglected, are minimally neglected in the czech republic. yet, and perhaps for this very reason, it is very important to address them. this article looks at the description of the different elements of the construction 4.0 concept and their possible application in transport, through the overall digitisation of the railways. keywords: construction 4.0, digitalisation, railway. 1. introduction the construction industry has innovative solutions and visions at its disposal. however, the key problem is that most construction companies do not consider themselves part of the evolutionary process and behave as if digital innovation has no real impact on them. this means that there is not yet a consensus among most construction executives about the reasons for change and therefore they are not adapting to it. the dominant uncertainties are productivity, costs, quality of work, workforce and skills management, and health and safety, which are just some of the opportunities that construction 4.0 could address. these need to be factored into a longer-term strategy to adapt in a way that will allow society to thrive in the next phase of digital evolution that will change many aspects of society [1]. as these innovative technologies make their way into different phases of the construction project lifecycle, there is growing concern about the future of jobs. while the use of technology, and robotics in particular, has the potential to increase productivity and safety, it should not necessarily reduce overall employment in the construction industry in the long term. it is expected that existing roles will evolve and new roles will emerge (e.g. employees with digital skills will be needed in addition to designers) [2]. construction 4.0 in the context of this article includes bim, virtual and augmented reality, 3d printers, advanced building materials, prefabricated and modular construction, big data and predictive analytics, 3d scanning and photogrammetry, autonomous vehicles and drones, cloud collaboration (real-time collaboration), the internet of things, or wireless monitoring and device connectivity [3]. construction 4.0 in railways is somewhat confronted with the fact that transport systems are complex in terms of technology and operation, and involve a range of human actors, organisations and technical solutions. for the operation and management of such complex environments, a viable solution is the use of intelligent computer systems, such as computerised traffic management systems for air traffic coordination or advanced monitoring and diagnostic systems in vehicles. in addition, the means of transport cannot compromise the safety of passengers by applying operational and maintenance activities. in fact, safety is becoming a more difficult objective to achieve using traditional maintenance strategies and computerised solutions are coming into the picture as the only option to deal with complex systems that interact with each other and try to balance the increase in technical complexity together with stable and acceptable reliability indicators [4]. the importance and significance of railway development through the construction 4.0 concept is evidenced by efforts to apply it abroad. these include: 1.1. smartrail 4.0: the swiss railway modernisation programme in the next few years, a number of technical production systems that are important for train travel will be replaced due to their age. at the same time, technologies are being developed at an ever-increasing pace, offering rail companies new opportunities to improve efficiency and punctuality [5]. sbb, bls, schweizerische südostbahn ag (sob), rhb, transport public fribourgeois (tpf) and the public transport union (ptu) would like to take advantage of this opportunity together. together, these companies are preparing the railway for the future as part of the smartrail 4.0 industrial programme. this is the first programme that covers the entire railway production, from timetable planning to train control [5]. as a first step, the concept of this overall architecture has been consolidated and refined in the period 19 https://doi.org/10.14311/app.2022.35.0019 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en pavel krupík acta polytechnica ctu proceedings up to 2020, and a more detailed roadmap has been drawn up. the smartrail 4.0 programme is divided into five sub-programmes and covers technological development for the next 20 years. the participating groups are modernising the rail system to ensure the robustness of future services, increase the capacity of existing infrastructure, improve the safety of rail workers and stabilise system costs [5]. 1.2. the construction 4.0 concept goes hand in hand with industry 4.0 thanks to digitisation, it is possible to use data obtained from traffic, for example, when planning the development of new railway lines. recent innovation projects announced by european rail operators deutsche bahn (db) and sncf are a perfect example of how railways are using frmcs to become a leader in industry 4.0. db, one of the world’s largest and most important railways, has announced that it will work with railway experts from nokia on the "digitale s-bahn hamburg" project, which will use 5g-based frmcs to deploy – as early as 2021 – highly automated trains and shunting operations on its s-bahn line in hamburg [6]. 2. aims the aim of this article is to describe the possibilities in terms of construction 4.0 in the planning and implementation of railways. therefore, building 4.0, in the concept of this paper, has the task of preparing the infrastructure to be ready for operation and maintenance. 3. methods the article mentions the possible use of construction 4.0 in the railway concept. the research methodology consists in searching the scopus open access database for relevant articles with innovative solutions and visions. furthermore, solutions that have been used in the recent past or are yet to be planned to be used in the future are also presented. 4. results and discussion 4.1. frmcs communication standard the latest standard for rail communications is moving from the widely adopted gsm-r to the frmcs (future railway mobile communications system) standard, with 5g as a key enabling technology. frmcs sets the stage for total connectivity and seamless endto-end operations with smarter automation and the ability to leverage new internet of things technologies for increased productivity and reliability breakthroughs. once frmcs is in place, railways can deliver – in the near future – key improvements that will increase operational speed and deliver a superior ride and travel experience for passengers and goods [6]. 4.2. human factor the human factor in innovation and development is in many cases one of the slowers of construction 4.0 transition and application. the natural human reaction to change is fear, rejection and even negativity. in the process of industry 4.0 development, it becomes therefore necessary and important to involve employees and students in these processes in an early and important way. their involvement can significantly accelerate the processes, so the development of it should take place in advance. similarly, the acquisition of technology knowledge or practical experience with machines, software or intelligent programs that they will encounter or actively use in the future to perform their work in the workplace will significantly accelerate and streamline adaptation, leading to increased time efficiency , reduced costs and time-consuming further training of employees [7]. the social (human) factor was identified as the most critical factor that has the greatest impact on the successful implementation of construction 4.0 in general, but other contributing factors (pestels: political, economic, social, technological, environmental, legal and security factors) suggest that these factors are interrelated and should be addressed simultaneously. the introduction of industry 4.0 in the construction sector through the principles of construction 4.0 would put the sector on a par with similar sectors such as manufacturing and automotive, where industry 4.0 is already widely used [8]. 4.3. bim currently, great efforts are being made to achieve a more general application of the bim methodology in the field of transport infrastructure, beyond specific elements such as tunnels, viaducts and bridges, as evidenced, for example, by the case study of the renewal of a railway line [9]. this highlighted the fact that in the field of transport infrastructure systems, the railway presents specificities and complexities that lend more interest and importance to the implementation of bim. the highest demands on railway safety (e.g. high-speed train) require more frequent monitoring and maintenance. in addition, the complexity of railway components and systems (infrastructure and vehicles) requires better tools for data analysis and processing. in this sense, bim will be a useful methodology with huge technical, economic and environmental benefits throughout the life cycle of railway infrastructure. one of the benefits of bim is the possibility to implement a track inspection system that includes all the usual geometric parameters in order to integrate maintenance planning linked to quality control tests. 4.4. artificial intelligence possible applications of artificial intelligence in building 4.0 are: 20 vol. 35/2022 construction 4.0 in the concept of the railway figure 1. railways 4.0 with 5g [6]. • design based on artificial intelligence for topology analysis of prefabricated elements – configuration, segmentation and optimization, • optimisation of supply chain management to reduce costs, • predicting various potential project risks based on historical and real-time data [10]. 4.5. worker 4.0 this is a concept that materializes the main principles and behaviors of workers in the construction 4.0 scenario. precisely because the productivity of the worker 4.0 movement is designed to manage and continuously improve processes, the four-stage pdca management method is appropriate for its implementation. first, an innovation project based on the worker 4.0 movement’s productivity methodologies for craftspeople needs to be planned. the next step is to complete data collection and processing. this means checking process performance to assess the level of mechanization. the last step is to act on productivity in the workplace, implementing control measures and setting up continuous improvement procedures. such an analysis should be based on the quantitative and qualitative values found in the measurements and the subsequent possibility of improving production indicators [11]. 5. conclusion the article dealt with the use of construction 4.0 elements that can be used in the construction and downstream operation of railways. some of these are (fortunately) very close to being used in practice. during the research it was found that one of the main factors why industry 4.0 is not as developed in the construction sector as in other sectors is the human factor, whether it is the reluctance to use innovation in practice, the lack of legislation or the need to train employees. the author believes that it makes sense to continue with the set direction of research in this area, but considers it appropriate to address other parts of the contemplated construction 4.0 in the railway sector than just bim. in this vein, the author welcomes the planned establishment of the "construction 4.0" platform, whose ambition is to bring together major research institutes, especially at technical universities, with progressive construction companies and to offer a space for mutual collaboration. the aim is to introduce new technologies and the use of industry 4.0 principles into construction and to increase the competitiveness of the czech republic in this area on a european scale. acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs20/100/ohk1/2t/11. references [1] r. woodhead, p. stephenson, d. morrey. digital construction: from point solutions to iot ecosystem. automation in construction [online]. 2018, 93, 35–46. issn 09265805, https://doi.org/10.1016/j.autcon.2018.05.004. [2] b. garcía de soto, i. agustín-juan, s. joss, j. hunhevicz. implications of construction 4.0 to the workforce and organizational structures. international journal of construction management [online]. 2019, 1-13. issn 1562-3599, https://doi.org/10.1080/15623599.2019.1616414. [3] p. krupík. systems for intelligent parking of trucks at motorway rest areas as part of construction 4.0. iop conference series: materials science and engineering [online]. 2021, 1203(3) [cit. 2022-01-18]. issn 1757-8981, https://doi.org/10.1088/1757-899x/1203/3/032052. [4] m. kans, d. galar, a. thaduri. maintenance 4.0 in railway transportation industry. proceedings of the 10th world congress on engineering asset management (wceam 2015) [online]. cham: springer international 21 https://doi.org/10.1016/j.autcon.2018.05.004 https://doi.org/10.1080/15623599.2019.1616414 https://doi.org/10.1088/1757-899x/1203/3/032052 pavel krupík acta polytechnica ctu proceedings publishing, 2016, 2016-03-26, s. 317–331. lecture notes in mechanical engineering. isbn 978-3-319-27062-3, https://doi.org/10.1007/978-3-319-27064-7_30. [5] smartrail 4.0: a modernization programme for the swiss rail sector. [online]. available online at: https://www.smartrail40.ch/index.asp?inc=en/ program.asp. [6] j. apel. railways set the pace for industry 4.0 transformation with frmcs/5g. nokia blog [online]. 2020. available online at: https://www.nokia.com/blog/railways-set_ -pace-industry-40-transformation-frmcs5g. [7] z. gerhátová, v. zitrický, v. klapita. industry 4.0 implementation options in railway transport. transportation research procedia [online]. 2021, 53, 23–30. issn 23521465, https://doi.org/10.1016/j.trpro.2021.02.003. [8] w. alaloul, m. liew, n. zawawi, i. kennedy. industrial revolution 4.0 in the construction industry: challenges and opportunities for stakeholders. ain shams engineering journal [online]. 2020, 11(1), 225–230. issn 20904479, https://doi.org/10.1016/j.asej.2019.08.010. [9] j. neves, z. sampaio, m. vilela. a case study of bim implementation in rail track rehabilitation. infrastructures [online]. 2019, 4(1). issn 2412-3811, https://doi.org/10.3390/infrastructures4010008. [10] m. wang, c. wang, s. sepasgozar, s. zlatanova. a systematic review of digital technology adoption in off-site construction: current status and future direction towards industry 4.0. buildings [online]. 2020, 10(11). issn 2075-5309, https://doi.org/10.3390/buildings10110204. [11] d. calvetti, p. mêda, m. chichorro gonçalves, h. sousa. worker 4.0: the future of sensored construction sites. buildings [online]. 2020, 10(10). issn 2075-5309, https://doi.org/10.3390/buildings10100169. 22 https://doi.org/10.1007/978-3-319-27064-7_30 https://www.smartrail40.ch/index.asp?inc=en/program.asp https://www.smartrail40.ch/index.asp?inc=en/program.asp https://www.nokia.com/blog/railways-set_-pace-industry-40-transformation-frmcs5g https://www.nokia.com/blog/railways-set_-pace-industry-40-transformation-frmcs5g https://doi.org/10.1016/j.trpro.2021.02.003 https://doi.org/10.1016/j.asej.2019.08.010 https://doi.org/10.3390/infrastructures4010008 https://doi.org/10.3390/buildings10110204 https://doi.org/10.3390/buildings10100169 acta polytechnica ctu proceedings 35:19–22, 2022 1 introduction 1.1 smartrail 4.0: the swiss railway modernisation programme 1.2 the construction 4.0 concept goes hand in hand with industry 4.0 2 aims 3 methods 4 results and discussion 4.1 frmcs communication standard 4.2 human factor 4.3 bim 4.4 artificial intelligence 4.5 worker 4.0 5 conclusion acknowledgements references https://doi.org/10.14311/app.2022.33.0467 acta polytechnica ctu proceedings 33:467–472, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague forced carbonation of recycled concrete aggregates marc-patrick pfleger∗, markus vill university of applied sciences vienna, favoritenstraße 226 a-1100 vienna, austria ∗ corresponding author: marc-patrick.pfleger@fh-campuswien.ac.at abstract. the production of concrete, especially the contained cement clinker, causes a high proportion of the manmade gaseous air pollution. studies show that the decarbonation of limestone effects roughly 8% of worldwide co2 emissions, which underlines the need for action and research. due to the present building and infrastructure stocks and their obsolescence, or rather replacement, the amount of demolition material is permanently increasing. in many cases the demolition material consists of high-quality concrete, which could be used again as recycled aggregate in the production process to conserve sources of raw materials. most of the concrete waste is not carbonated which means that the potential as co2 sink can be exploited in the course of the recycling process. since the recycled and carbonated material is used as a replacement for primary raw material, the state of carbonation itself is irrelevant to any steel corrosion or associated problems. the main objective of this work was to investigate the co2 absorption of concrete waste, to investigate and prove its use as a possible carbon sink. at this point, the carbonation velocity had to be optimised by the use of special equipment. first test series show that the ecological footprint of concrete can be significantly reduced if concrete waste is used for carbon storage and admixed as recycled aggregate during the production process. the aim of this study was to point out further possibilities for the ecological optimisation of concrete besides alternative binders, which are the subject of research in various projects worldwide. keywords: carbonation, concrete, recycled aggregates. 1. introduction the protection of the environment and the global climate is currently a major field of action. not only changing climate conditions are perceptible, but also rising raw-material demand, land consumption and the pollution of important habitats occur as main problems. since every industrial sector has more or less impact on the environment, the question arises where significant entry points for ecological optimisations can be found. often a close connection can be found between the amount of materials used by a particular business branch and its environmental impact. the construction industry and related trades are associated with an enormous need for raw materials. further processing uses high amounts of energy, which means that there is a tangible potential for greenhouse gas reduction. one of the most commonly used materials for civil engineering structures and buildings at present is concrete or rather reinforced concrete. especially the contained cement clinker is a highly climate relevant substance. according to the current state-ofthe-art, structural concrete cannot be adequately replaced, neither technically nor economically. its robustness, stability and formability, in combination with low material and manufacturing costs, are the main reasons for the continuously rising demand. air-pollution is increasing analogously to the growing amounts of produced binder or rather the contained cement clinker. the fact that nowadays about 8% of the worldwide co2 emissions (by population) result from producing cement has to be highlighted as absolutely critical. the mentioned emissions result only partly from combustion processes to provide the needed heat to process clinker out of lime stone. if modern technology is used, over 50% of the overall co2 emissions concerning the cementitious binder production result from the lime stone decarbonation, as seen in table 1. emissions due to the raw material decarbonation are process-related and cannot be substituted, if the known characteristics of cement shall not be modified. especially the corrosion protection of the reinforcement depends on the alkalinity of the hardened concrete. this paper brings the described problems with greenhouse gas in context with rising amounts of demolition materials, especially concrete recycling products. the majority of concrete waste is not carbonated, in particular the inner sections of structural elements of many civil engineering structures, like bridges with large thicknesses of concrete members. 2. background for ecological optimisation of concrete as described before, concrete is an indispensable construction material, because of its technical charac467 https://doi.org/10.14311/app.2022.33.0467 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en marc-patrick pfleger, markus vill acta polytechnica ctu proceedings raw material decarbonation combustion of fossil fuels combustion of biogenic fuels total 2016 524 260 72 856 2017 517 256 71 844 2018 515 254 75 844 table 1. specific co2 emissions in [kg] per [to] clinker in austria [1]. teristics. therefore, optimisations concerning the ecological footprint of concrete can actually only be made if the occurring of actual disadvantages can be consistently avoided. in general, improvement potential can be seen in alternative binder recipes or using high-grade recycling processes for an ideal reuse of concrete waste. the minimum share of cement clinker in the most commonly used standardised cement mixtures (cem ii) is about 65%. the rest consists, depending on the exact cement type, of fly ashes, blast furnace slag or lime dust. the currently available amount of fly ash and slag is already completely used up, so that no more substitution can be expected. furthermore, due to the focus of future energy policy, the available quantity of additives, based on fossil combustion processes, will decline. therefore an extensive amount of research activity is done worldwide to develop alternative binders for concrete applications, like [2], [3]. commonly the main objective is to reduce or avoid the share of cement clinker in the binder mixture. problems with durability may occur as a consequence and are not investigated enough in long term tests. apart from binder optimisation the rising stream of demolition material has to be involved in high quality recycling. particularly in urban regions new buildings can only be produced if a site is previously cleared from its original building stock. it can be observed worldwide that a building’s life cycle is decreasing, which means that the building stock becomes a more important source of raw material in the sense of "urban mining". generally demolished concrete structures are processed by one or more crushers to receive a useable concrete granulate. depending on the valid standards, this concrete recycling aggregate can be used with a certain percentage as replacement for natural aggregates. concerning the national documents of austria, additional differentiation is made for admixing, depending on the grain size of the recycling product [4]. usually those recycled aggregates have a grain size distribution from 0 to 30 mm or smaller. the maximum grain size in general depends on the desired configuration of the jaw crusher. 3. carbonation of concrete 3.1. carbonation as concrete damage the carbonation of concrete is usually known as a classic concrete damage and can be described as the reaction of co2 from the surrounding atmosphere with calcium hydroxide, as a product of the hydration of cement clinker. the calcium hydroxide is contained in the pore water of a concrete matrix and its concentration may vary due to the particular binder mixture and the surrounding conditions during hydration. the specific concentration of lye determines the alkalinity of the pore water and thus the capability to prevent the reinforcement from corrosion. under the absorption of co2 this potential declines over time, starting from the concrete surface. therefore the tolerable carbonation depth is generally limited to the minimum concrete cover of the reinforcement. concrete technological measures are taken to prevent fast carbonation. main ambient parameters for the carbonation velocity can be: • carbon dioxide level of the surrounding atmosphere, • ambient air and concrete temperature, • concrete moisture, • pressure. additional influencing factors result mostly from the material itself such as content of cement, porosity or (micro) cracks which provide an increased surface area for carbon dioxide absorption and reaction [4]. 3.2. carbonation of concrete as carbon sink the inner cross-section parts, or rather the area between the outer reinforcement layers, are usually not carbonated until the demolition of a structure. emitted co2 caused by the calcination of the binder remains in the atmosphere and has a negative impact on the materials ecological footprint. the question arises as to which surrounding conditions and related time span are required until the carbonation process is completed after a concrete structure is recycled into aggregates. as described before, the carbonation of concrete is a natural process, though natural conditions only lead to a slow reaction process with the actual co2 consumption not defined. nevertheless, there is carbon absorption at any time where the essential conditions, such as concrete moisture, are given. the chemical conversion of calcium hydroxide into calcium carbonate effects that a solution transforms into solid [5]. as a result, the gas permeability of the cement matrix decreases and 468 vol. 33/2022 forced carbonation of recycled concrete aggregates figure 1. not-carbonated share of a cross-section. the carbonation velocity asymptotically approaches zero within the grain [6]. based on this knowledge, an artificial method shall be preferred in comparison to carbonation under natural conditions. controllable parameters and the possibility of measurements during the carbonation ensure a consistent recycling product [7]. 4. experimental setup and methods the test setup was designed to investigate forced carbonation of recycled concrete aggregates, which means that no correlation with natural carbonation could be expected. the first step of the experimental carbonation test series was to determine the exact amount of carbon dioxide which can be bound by a concrete matrix. for this purpose, special equipment was necessary, since standard carbonation chambers do not allow to vary and monitor the reaction conditions in real time as desired [8]. especially the used experimental setup allowed to monitor and control the exact co2 consumption of the sample material in real time. further it was planned to find the optimal parameters to fully carbonate a particular sample as fast as possible. 4.1. reaction parameters as basis for the equipment design, the previously defined external parameters were taken into consideration. thus temperature, humidity as well as the amount of co2 in the sample atmosphere should be varied. increasing the ambient pressure was planned for future sample series. the range of reaction parameters was defined according to feasible values in large-scale techniques or rather concerning industrial facilities as shown in table 2. from to co2 0,04% 30% humidity 0% 100% temperature 20◦c 65◦c table 2. range of reaction conditions. 4.2. equipment design the equipment consists of a carbonation chamber that can be top-loaded with the sample material. a removable grate is filled with the sample aggregates up to 25cm layer thickness. the container sits hermetically on a catch, so that a forced flow of the ambient atmosphere through the material can be ensured by the integrated fans. the perfusion of the sample container happens from bottom to top. as shown in figure 2, all fittings and connections are located at the cover of the carbonation chamber. so only the tempered and humidified granulates are inside the reaction tank which consists of stainless steel. the acryl glass lid is prepared for an attachable steel frame to ensure sufficient stiffness if tests with overpressure will be carried out in the future. a main requirement was that the co2 absorption of the samples should be measurable at any time of the carbonation process. a standard carbonation chamber does not allow a real-time monitoring of the actual co2 consumption as desired. further, any implementation of additional custom equipment is easier, dealing with a setup from scratch. it is planned to generate a completely automated process by using artificial intelligence to set the ideal reaction conditions and vary them over the test’s duration. due to the moistened material and the increased humidity inside the chamber it is not possible to obtain reasonable results by measuring the sample’s mass during the experiment. the material moisture varies during the test. additionally, condensation will distort 469 marc-patrick pfleger, markus vill acta polytechnica ctu proceedings figure 2. simplified scheme of the test set-up. any weighing process inside the chamber. co2 reacts with the calcium hydroxide contained in any concrete waste to caco3, so that a direct conclusion can be drawn about the absorbed carbon. therefore the increasing weight of the aggregates shall be proved by other boundary parameters such as the decreasing gas bottle mass, the opening time of the valve and the flow volume of co2. all these three parameters where monitored and compared to validate the accuracy of the data. the measured values are documented and analysed in real time, which allows statements about the actual co2 consumption according to current reaction conditions. at the end of the experiment dry mass of the sample can be weighed and compared with the initial mass in order to verify the results derived from the experiment. data-logging and the real-time analysis is physically located in an extra housing where also the co2 probe was placed. gas from the reaction chamber is extracted, analysed and pumped back. since only carbon consumption was expected here, no precautions were taken to reduce the carbon dioxide share of the test atmosphere. due to the more or less constant reaction of co2 the valve is only opened when necessary to keep the concentration just as desired. ideally the reaction conditions are varied during the carbonation process to achieve more efficiency. furthermore, initial tests show that varying the reaction conditions according to different cement types gives better results. 4.3. samples granulated concrete with accurately known recipe was found to be a suitable sample material for the first test series, which also should serve for data verification. several concrete test cubes with 15cm edge length were produced and stored in the air for 60 days under standardised conditions for 60 days exposed to air. 5. test series 5.1. sample concrete aggregates all experiments carried out with crushed concrete aggregates were done by using the same concrete mixture to achieve ideal comparability. the determined samples were made from c25/30 containing cem ii b with a maximum grain size with 22mm in the matrix. the binder portion was defined as 15% of the wet concrete mass. the water-cement ratio was kept at 0.5. 5.2. preparation of the sample material the first step of each test series was the determination of the samples dry mass. to extract residual moisture the grains were put in a dry kiln until the dry weight was proven by mass stability. as basis for all test series the aggregate mixture was separated into grain fractions 0 to < 8mm and 8 to 22mm. the fraction <8mm had 26.5% of the total mass, thus 73.5% were weighed with 8-22mm. depending on the particular experiment the samples were moistened with a defined amount of water to exploit the influence of material moisture to the carbonation velocity. in fact, the ideal material moisture cannot be expressed with a simple percentage, since the porosity, in particular the pore diameter, is a key factor for the amount of water needed to moisten the pore walls [5]. 470 vol. 33/2022 forced carbonation of recycled concrete aggregates figure 3. carbonation results for c25/30, cem iib, grain size 0 − 8mm. 5.3. structure of the grain fractions after separating the grain fractions visual inspection alone showed that most of the cementitious parts of the crushed down concrete find themselves mainly in small grains. this means that also most of the potential carbon sink was initially expected in small grain fractions. 5.4. influence of the layer thickness several carbonation tests were done with mainly small grain sizes from 0 to 8mm. as expected, the permeability of a sample layer consisting of only small grain fractions is poor. the effect is seen when checking the carbonation of grains from the centre of the bulk. when testing coarse aggregates, a more consistent carbonation through the whole layer depth was verified, whereas aggregate mixtures do not ensure enough permeability as well as recycling sands. the issue with poor permeability can be obviated by using a non-static process. 6. results 6.1. concrete aggregates the following results show the total amount of bound carbon dioxide in the samples over the test duration. ideally the mass increasing shall not be compared directly with the total mass at the experiment’s start, because of the contained natural aggregates in the concrete matrix, which are not reactive. when the ecological improvement is calculated, only the binder or cement clinker share shall be considered. 6.1.1. grain sizes under 8mm figure 3 shows exemplarily how carbon absorption proceeds at grain fractions below 8mm. the experiments duration was 24 hours with 20% co2 in the carbonation chamber, heated to 40◦c. the carbon dioxide concentration as well as the temperature was constant during the whole test. the sample’s dry mass was determined to 11.289g at the beginning. it can be seen that the co2 consumption is very high at the beginning, which can be explained by the rapid carbonation of the surface areas of the crushed figure 4. carbonation results for c25/30, cem iib, grain size 8 − 22mm. down concrete and the high share of cement of the recycling sands. the further course of the graph varies visibly over time. in this case this phenomenon cannot be explained by the saturation of the aggregate. when the test was completed it was seen that poor carbonation results were achieved in the centre of the bulk. this can be explained by too little gaseous exchange due to the high packing of the small grains. the test ends with 238g co2 bound in the sample, resulting in 21g co2 per 1kg recycled aggregate (grain size < 8mm). 6.1.2. grain sizes over 8mm figure 4 shows exemplarily how carbon absorption proceeds at grain fractions over 8mm. the outer parameters of the experiment were exact the same as mentioned in 6.1.1., which means: 24 hours of carbonation with 20% co2 at 40řc in the test setup. both the carbon dioxide concentration and the temperature were stable during the whole test. the sample’s dry mass was determined to 11.385g at the start. again it can be seen that co2 consumption is high at the beginning, whereas the gradient is less prominent when smaller fractions are analysed. the graph shows a clearly more flat carbonation curve at the start due to the smaller surface to volume ratio. furthermore, the reaction velocity decreases permanently over time and approaches zero. after the removal of the sample material it was seen that constant carbonation took place in all layers of the bulk. the test ends with 284g co2 bound in the sample, resulting in 25g co2 per 1kg recycled aggregate (grain size 8-22mm). the test series show that co2 can be absorbed with 2.1% to 2.5% of the sample mass in a static process. that mass increase is accompanied by the formation of lime in the concrete matrix which means that simultaneously the porosity of the samples must decrease. this phenomenon was proved by an additional saturation experiment before and after the carbonation process. 471 marc-patrick pfleger, markus vill acta polytechnica ctu proceedings 6.2. exemplary presentation of the ecological benefit based on the results from 6.1. the ecological advantage of recycling concrete with forced carbonated can be pointed out. due to current austrian national standards, the admixture of up to 50% recycled which is explained by unpredictable increased water needs [4]. the following example states, that 50% of the natural aggregates are substituted by a forced carbonated recycling product. it is expected that a concrete strength of c25/30 with cem ii shall be achieved. it can be assumed that 900kg of carbonated aggregates are admixed while concrete production. thereby 22.5kg carbon dioxide can be credited per m3 fresh concrete if the ecological footprint of the material is calculated. in comparison to the carbon emissions due to the binder production with approximately 170kg per m3, an improvement of about 13% can be achieved. 7. conclusion the exemplary stated test series show the potential of forced carbonation of recycled concrete aggregates in principle. an absorption of 2.1% to 2.5% of the sample’s mass was achieved by the test setup. as the cement clinker production increases permanently worldwide and thereby the carbon dioxide emissions rise, ways for ecological optimisation must be found. the experiments show that forced carbonation is suitable to utilise existing carbon sink potentials of demolished concrete. if the carbonation of recycled concrete aggregates can be implemented in current processes with high carbon dioxide emissions, such as cement plants, the ecological footprint of concrete actually can be reduced. currently, the life-cycle assessment of concrete does not take recarbonation into account. in practice carbonation of structural concrete elements is generally present. however, only the surface areas can be practically considered during a buildings period of use. at the end of a structure’s life-cycle, a demolished concrete is stored in piles and may be used as recycled aggregates for mixing concrete or technical backfill material. as the test showed an increased natural carbonation cannot be expected, because of insufficient air flow or rather gas exchange through the stored material. a forced carbonation process can efficiently use the crushed down material as carbon sink in a much more predictable period of time. for this reason an uninterrupted recycling process leading to the final product is possible. the further purpose is to show if the expected differences in the carbonation behaviour of different cement types will occur during the tests. if decisive differences are seen, the findings can simplify a possible automation for the adjustment of ideal carbonation parameters. references [1] g. mauschitz. emmissionen aus anlagen der österreichischen zementindustrie, berichtsjahr 2018. technical university of vienna, 2019. https://www.zement.at/downloads/downloads_2019/ emissionen_2018.pdf. [2] t. higuchi, m. morioka, i. yoshioka, et al. development of a new ecological concrete with co2 emissions below zero. construction and building materials 67:338-43, 2014. https: //doi.org/10.1016/j.conbuildmat.2014.01.029. [3] j. juhart, c. baldermann, g-a. david, et al. erescon neue betone für den infrastruktur-bereich, energieund ressourcenoptimiert. eigen, 2015. https://www2.ffg.at/verkehr/studien.php?id= 1112&lang=de&browse=program. [4] önorm b 4710 1:2017-06-15. austrian standards institute, vienna, 2017. [5] h. grube, j. krell. zur bestimmung der carbonatisierungstiefe von mörtel und beton. beton 36(3), 1986. [6] j. gajda, f. m. miller. concrete as a sink for atmospheric carbon dioxide: a literature review and estimation of co2 absorption, portland cement concrete: portland cement association, 2000. . [7] c. gehlen. probabilistische lebensdauerbemessung von stahlbetonbauwerken. dafstb (= dafstb-heft. band 510). beuth, 2000. [8] m. seidemann, a. müller a, h. ludwig. weiterentwicklung der karbonatisierung von rezyklierten zuschlägen aus altbeton (2. phase: prozessoptimierung im labormaßstab und technologieentwurf), 2015. https://www.dbu.de/opa c/ab/dbu-abschlussbericht-az-23689_02.pdf. 472 https://www.zement.at/downloads/downloads_2019/emissionen_2018.pdf https://doi.org/10.1016/j.conbuildmat.2014.01.029 https://www2.ffg.at/verkehr/studien.php?id=1112&lang=de&browse=program https://www.dbu.de/opac/ab/dbu-abschlussbericht-az-23689_02.pdf https://doi.org/10.14311/app.2022.33.0552 acta polytechnica ctu proceedings 33:552–557, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague methodology for evaluation of the life cycle of buildings with a focus on the private sector renáta schneiderová heralová∗, eduard hromada, jakub holcman, stanislav vitásek czech technical university in prague, faculty of civil engineering, department of construction management and economics, thakurova 7, 166 29 praha 6, czech republic ∗ corresponding author: heralova@fsv.cvut.cz abstract. the paper describes the innovative methodology for the evaluation of the life cycle of buildings, which will be used in construction projects of skanska, or other construction companies, in selecting the optimal solution from the point of view of construction engineering and lcc. the aim is to help lcc to be perceived by construction practice as a topic that needs to be addressed at the beginning of each construction project. the quantification of life cycle costs based on relevant input data on the technical parameters of the construction, structural elements and equipment, the time period of incurring the related costs should be an important basis for the decision of the investor, designer and future user of the building also with regard to environmental and social aspects, and in particular the long-term economic consequences. keywords: building, life cycle, life cycle costing, methodology. 1. introduction this paper deals with innovative methodology in construction company skanska concerning the evaluation of construction’s life cycle costs (lcc). mentioned methodology will be applied to construction projects of blocks. this will lead to optimal choices accounting for technical aspects of construction, and lcc from the very moment of designing a construction. according to the international standard [1], life cycle costing is a valuable technique that is used for predicting and assessing the cost performance of buildings. life cycle costing is one form of analysis for determining whether a project meets the clients performance requirements. life cycle costs (lcc) mean involving all of the costs deriving from the use of buildings during their entire life span. the lcc methodology is a tool for evaluating these costs over time. its main goal is to assess the various designs, where those designs differ not only in their acquisition costs but also in their operational and maintenance costs. assessment of life cycle costs is a valuable tool for both the construction company and the eventual user of construction. both company and end user are clearly informed about total costs from the beginning of construction. therefore, they do not need to concentrate only on investment costs or use them as the single criterion of choice. options of construction may be designed as soon as the designing phase of a project. so the end user obtains higher added value and may compare variant designs to reach optimal value for money. variant designs enable better choices of construction suitable for the end user’s requirements. end user may plan eventual maintenance costs accordingly. this advantage may be the key factor especially for end users who finance the construction with third party resources, such as loans. life cycle planning lowers the amount of risk in the investment. transparency can be reached as to operational costs of construction. with life cycle costing it will be possible to benchmark competing projects. lcc analysis has already been implemented as one of the measures for varied certifications, which serve for project evaluation (breeam, leed, sb tool, etc.). 2. literature overview buildings and constructions are long-term possessions. this is the reason why all decisions connected with a construction project have long-term and significant impact [2]. construction project investors have often focused only on the acquisition costs, when they were about to make decisions on such matters as the building design, facilities, fittings. they often neglected future operation and maintenance costs [3]. life cycle costing is a method assisting an effort to estimate the total cost of ownership. the method is able to help in making decisions within building investment projects [4]. life cycle costing is particularly useful for estimating total costs at the early stage of a project [5]. life cycle costing consists of all costs directly connected to a construction during its whole existence. that is building costs, maintenance and renovation of construction units’ costs, facility equipment costs, facility management and operation costs, demolition costs. [6] many authors also link other factors closely with life cycle costing. among such factors, let us name externalities, non-construction costs and income which arise from 552 https://doi.org/10.14311/app.2022.33.0552 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 lcc methodology for private construction sector id category designation description 1 const. mcw1 works and provisions 2 const. 1 ground works 3 const. 2 ground solidification 4 const. 3 vertical and complete constructions 5 const. 4 horizontal constructions 6 const. 6 finishes and flooring 7 const. 60 finishes internal 8 const. 63 flooring and floor construction 9 const. 9 other constructions and works, demolition 10 const. 90 other constructions and works 11 const. 94 scaffolding, construction lifts 12 const. 96 demolition of constructions 13 const. acw2 works and provisions 14 const. 71 insulation 15 const. 711 water, moisture and gas insulation 16 const. 76 constructions 17 const. 763 assembled constructions 18 const. 764 external plumbing 19 const. 766 carpentry 20 const. 767 ironworks 21 const. 768 panel work windows 22 const. 77 flooring 23 const. 771 floors tiled 24 const. 775 floors wooden, frieze, floating 25 const. 776 floors thin finish 26 const. 777 floors cast, synthetic 27 const. 778 floors double 28 const. 78 finishing works 29 const. 781 ceramic tiles 30 const. 784 paintwork and wallpapers 31 const. 788 fire-fighting devices 32 const. 789 facade casing and contact insulation 33 const. 790 roof constructions 34 const. 791 roof vegetational 35 const. 792 roof trafficable 36 const. 793 roof non-trafficable 37 const. 794 roof movable 38 monolith mcw1 works and provisions 39 monolith 2 ground solidification 40 monolith 27 foundations 41 monolith 3 vertical and complete constructions 42 monolith 31 construction walls 43 monolith 33 columns, pier, framework 44 monolith 34 walls and partitions 45 monolith 4 horizontal constructions 46 monolith 41 ceilings and ceiling constructions of buildings 47 monolith 43 stairway constructions and ramps 48 infrastructure 101 sewage connection 49 infrastructure 102 waterway connection 50 infrastructure 103 heating connection 51 infrastructure 105 gas connection 52 infrastructure 106 weak current service line 53 infrastructure 107 outdoor water distribution 553 r. schneiderová heralová, e. hromada, j. holcman, s. vitásek acta polytechnica ctu proceedings 54 infrastructure 108 outdoor sewage 55 infrastructure 109 outdoor gas distribution 56 infrastructure 110 outdoor weak current service line 57 infrastructure 111 outdoor heavy current service line 58 infrastructure 112 outdoor lights 59 infrastructure 113 electricity substation 60 infrastructure 114 ground preparation 61 infrastructure 115 roads, pavements, public equipment 62 infrastructure 117 landscaping 63 infrastructure 118 fencing 64 infrastructure 119 traffic signs 65 infrastructure 120 structural walls 66 air conditioning 240 air conditioning 67 heating / cooling 731 central heating 68 sanitary units / gas conduit 721 sanitary units 69 sanitary units / gas conduit 723 indoor gas conduit 70 heavy current 210 heavy current, lightning rod 71 weak current 211 weak current 72 electric fire alarm 212 electric fire alarm 73 measuring and regulation 213 measuring and regulation 74 fire extinguishing equipment 214 semi-stable fire extinguishing device 75 lifts 330 lifts, escalators 1 major construction works, 2 additional construction works. table 1. structure of construction units and technologies with regard to lcc. the use of a construction in a specific location [7]. other authors are dealing with the topic of nearly zero energy buildings (nzeb) and cost-optimal level [8]. building certification is also closely related to the issues addressed. there is a national certification system sbtoolcz in the czech republic. the certification systems were focused on new construction originally [9]. 3. construction project database the lcc methodology will include a draft of skanska construction project database. the database has already been created with microsoft access. it currently contains rough data of projects. the database will serve as an individual tool as well as data source for other programs. for the sake of clarity, the table 2 shows the structure of construction units and technologies with regard to lcc. decisions concerning the choice of the construction technology and construction materials are not any longer based entirely on technical and economic attitudes, but are becoming increasingly influenced by life cycle cost and environmental considerations. in fact, the capability to influence the outcomes of whole life ownership is enormous during the design phase. the types of material specified, the quality of the design and the contracting method have to be chosen directly upon operation and maintenance costs. for example, choice of materials for horizontal and vertical supporting structures influences the total building time, transportation of materials and construction life span. moreover, it influences the environmental impact of construction. it is well documented that the majority of decisions about operating, maintenance and rehabilitation costs are predetermined at the design stage. basically, it is crucial to establish a device at the design stage that brings together the life cycle costing, service life, environmental life cycle assessment, and risk associated with decisions taken at this stage. the proposed methodology will enable variant solutions of a construction project which will help the construction company minimize environmental impact and life cycle cost. regarding the construction site itself, these may include air pollution caused by transportation and dust fall. the level of dust in the air may be reduced by sprinkling dirt roads within construction sites. however, water costs for sprinkling enter the total building costs, and they are not perceived as either economically efficient or ecological in terms of introducing water to the landscape in appropriate ways. moreover, this example is set in the operational phase which is currently under minimal surveillance when investment decisions are made. the methodology assigns economical life cycle of a construction to be 30 40 years. within this period it is not necessary to renovate carry constructions or stabilize them in any way unless an emergency situation. most construction elements are protected by finishes which need maintenance. this will be taken into account in the application. therefore, carry elements do not appear in the following phases of the life cycle, unlike maintenance and renovation costs of 554 vol. 33/2022 lcc methodology for private construction sector maintenance level (lm) annual costs service life sub-standard maintenance reduction by 20 % reduction by 5 % standard maintenance no change no change above-standard maintenance increase by 20 % increase by 5 % table 2. maintenance level. other construction units. (demolition of construction is not considered in the model.) concrete constructions may only enter the maintenance phase in case there are face concrete units which would need renovation during the building life cycle. these issues can be treated as soon as the design phase the relation of construction demands in operational and maintenance phase. use of concrete suggests variants of monolithic constructions or prefabricated units. monolithic ceiling construction has ultimately different characteristics than prefabricated construction. with prefabricated units, there is higher risk of plaster ruptures or there are higher costs of plaster in such quality as to ensure consistent finish. when considering the eventual life cycle phase demolition waste management would need to be taken into account, as well as transportation and/or recycling materials (such as concrete). environmental issues and health protection should not be underestimated in this phase, as well as costs connected with these issues. an environmental lcc methodology takes into consideration not only the main cost categories (investment costs, operating costs, maintenance cost), but also external environmental costs. the environmental impact may carry consequential costs for society. in general, constructed buildings, materials and products may have environmental impacts. the processes of manufacture, transport, assembly/disassembly, maintenance and disposal linked to buildings result, for example, in emission of greenhouse gases. subsequently, significant investment will be needed to neutralize these consequences in the future. calculation should include the costs of greenhouse gas emissions calculated using cumulated carbon costs for the period which is analyzed (through prices of emission allowances). the lcc methodology evaluates costs of mitigating/reducing environmental impacts. as a result, the best value solution could be identified from both economic and environmental perspectives. 4. lcc calculation tool life cycle cost calculation will be performed by a tool which was created by the authors of this paper. the application sw is used for the calculation of the lcc for ground structures in the microsoft excel operating environment. the lcc calculation includes the costs of acquisition, maintenance, service, renovation and energy consumption for designed development project options. it is a complex tool the operation of which is subject to the entry of input values to several separated topic-related spreadsheets. this separation was chosen for the sake of better clarity of data entered to the sw. the tool offers the possibility of the entry of three options of the design. each option makes it possible to follow up to sixty years period of the project including the possibility to postpone the starting date of its operation. this time shift takes into account a postponement and the lead time of the construction works. the following sections describe individual spreadsheets that are included in the calculation of the lcc in the designed application sw. the spreadsheet recapitulation is used to summarize the basic identifiers of the project and clear evaluation of individual lcc calculation options. the software shows both figures and graphic representation of individual lcc sections and the total lcc of the project including key parameters of the calculation (the nominal discount rate and the monitoring period). the lcc formula components are converted to the current value in individual periods using the discount factor. the additional (comparative) indices include "lcc per 1 m2 of the floor area", and "the ratio (budget costs/lcc)". the spreadsheet acquisition costs is used for the calculation of the total acquisition construction costs of individual project options that are divided into two groups. the first category includes the structuresť acquisition costs included in the maintenance, operation and renovation. the other category includes structures where the maintenance, service and renovation costs are expressed by a lump sum. in contrast to the above, costs of structures included to the maintenance, operation and renovation are calculated on individual basis. the model parameters spreadsheet is used for the entry of key inputs for the calculation of the lcc total net present value such as the monitoring period, inflation rate, actual discount rate, anticipated postponement of the project implementation and anticipated construction period for designed options. the spreadsheet structural element is used for the entry of key inputs for the calculation of the total present net value of the lcc such as the acquisition price, area, service life, maintenance costs, individual structural element service regularity and costs. moreover, the spread sheet makes it possible to enter the maintenance level (lm) to the structural elements. it uses a specific coefficient to modify their maintenance costs and their service life (for details see the table 2 below starting values). 555 r. schneiderová heralová, e. hromada, j. holcman, s. vitásek acta polytechnica ctu proceedings figure 1. evaluation of design options life cycle cost. the spreadsheet operation_energy is used for the calculation of energy price during the monitoring period for individual design options. the spreadsheet includes respective results diagrams for a better presentation. three energy types consumption is monitored: electric power, gas and water. the calculations are based on designed energy consumption for individual technological systems included in the project (usually in the document called building energy performance certificate (bepc). each option specifies eight pre-defined technologies that could be included in the project. moreover, there is a so called optional item available that can be used for the entry of any technology whatsoever that has been added to the project. not all technologies have to be accounted for in all cases. sometimes only the most relevant ones from the point of view costs can be included. on the spreadsheet structure, there is a key calculation tool for the partial parameters of the total net present value of the lcc such as the service life including maintenance level, number of renovation cycles in the monitoring period, acquisition costs, renovation costs, maintenance costs including the maintenance level, the number of service cycles in the monitoring period and costs of service of structural elements of individual design options. the spreadsheet renovation, maintenance and operation is primarily designated for the graphic representation of acquisition costs, renovation costs, maintenance costs, service costs and operation costs broken down to individual years converted to the present value using the discount factor. a summary table is available for the service, renovation and maintenance with three maximum and minimum cost values for individual structural elements. 5. case study a benefit of the lcc calculation tool is the possibility to model the life cycle costs for individual design options with different input parameters. that means that the user can choose the more efficient solution for him/her. for instance, he/she may find out what will be the impacts of a more expensive but energy saving option like not only on energy costs but on the total lcc, too. similarly, if he/she includes "maintenance" free equipment that requires higher acquisition costs, there is an apparent impact residing in the reduction 556 vol. 33/2022 lcc methodology for private construction sector of the maintenance costs but also in the value of the total costs of this project design option. the model includes a graphic processing of data for the presentation of results to all stakeholders. the data recapitulation is the most useful tool for the analysis of the building lcc for any managerial decisions. this is the presentation of results and basis source document for the adoption of respective decisions in the investment project preparation phase. any project can be integrated to the calculation model. having added various design options, we can find out the impact of the given change on the total costs. anticipated changes can be addressed on individual basis or as summary change sets. the given recapitulation example clearly indicate the load of the project when we take into account the maintenance, service and building operation (energy) costs. the key importance of the developed model resided in its capability to define the advantageous combination of the designed technologies for the economical operation. this is one of the most essential aspects from the point of view of the building life cycle costs. the optimal combination of the technological units becomes an essential quantity for the entire building life cycle. the model cannot secure the presence of the complete offer of technologies as may be available on the construction market. properly specialized and authorized designers should be aware of such information. the model can look at the technologies from the framework of the interval of performance, consumption of energy or operating costs. the combination of the most expensive systems does not guarantee the problem-free and low cost operation of the building. chosen technologies have to be able to communicate with each other, i.e. they have to be effective. this is the most demanding part of the lcc calculation. the developed tool is therefore designed as an aid for those company workers who are in charge of the preparatory phase of the project. given the fast development of technologies, it will be necessary to see the future results as just recommending but not binding data that are correct for the entire economic life period. 6. conclusions one of tools use for the assessment of the economic sustainability of designed buildings is the analysis of life cycle costs that is executed based on relevant input data on technical parameters of the building design, structural elements and equipment, related time interval of the costs that incurred. the analysis becomes an important underlying document for the decision of the investor on the identification of the best option of the building design taking into account also environmental aspects and long-term economic consequences. the analysis may be a significant marketing tool focused on the future building user to whom it presents the environment-friendly parameters and economic sustainability of the building. the evaluation of a construction project on the basis of its life cycle is included in certification systems evaluating the quality of buildings. the analysis of life cycle cost is assessed as a pertinent tool for improving the sustainability of building structures. it is advisable that both professional and lay population realize the necessity of exploring both acquisition costs and all combined other costs during the entire life cycle when making decisions about new constructions. acknowledgements this paper was supported by the technology agency of the czech republic, project id tn01000056, title "centre of advanced materials and efficient buildings cameb". references [1] iso 15686-5:2017(e). buildings and constructed assets service life planning part 5: life-cycle costing. international organization for standardization, 2017. [2] m. ryghaug, k. h. sørensen. how energy efficiency fails in the building industry. energy policy 37(3):984-91, 2009. https://doi.org/10.1016/j.enpol.2008.11.001. [3] m. jakob. marginal costs and co-benefits of energy efficiency investments. energy policy 34(2):172-87, 2006. https://doi.org/10.1016/j.enpol.2004.08.039. [4] r. flanagan, a. kendell, g. norman, et al. life cycle costing and risk management. construction management and economics 5(4):s53-s71, 1987. https://doi.org/10.1080/01446193.1987.10462093. [5] u. bogenstatter. prediction and optimization of life-cycle costs in early design. build. res. information 28(5-6):376-86, 2000. [6] r. heralova building’s value assessment using the utility and the lcc. cesb 2007. prague, 1:126-131, 2007. [7] x. meng, f. harshaw. the application of whole life costing in pfi/ppp projects. proceedings 29th arcom 2013, united kingdom. pp. 769-778, 2013. [8] j. karasek, j. veleba j. development of nearly zero energy buildings and application of cost optimum. business & it, vol. vii(2), pp. 18-25, 2017. doi:https://doi.org/10.14311/bit.2017.02.03. [9] d. macek. criteria for national rating system for buildings sbtoolcz. business & it, vol. vi(2), pp. 2-9, 2016. doi:https://doi.org/10.14311/bit.2016.02.01. 557 https://doi.org/10.1016/j.enpol.2008.11.001 https://doi.org/10.1016/j.enpol.2004.08.039 https://doi.org/10.1080/01446193.1987.10462093 https://doi.org/10.14311/bit.2017.02.03 https://doi.org/10.14311/bit.2016.02.01 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0269 acta polytechnica ctu proceedings 38:269–274, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague measurement and cfd analysis of a local radiant cooling solution shiva najaf khosravi∗, helene teufl, ardeshir mahdavi tu wien, department of building physics and building ecology, karlsplatz 13, 1040 vienna, austria ∗ corresponding author: shiva.nkhosravi@gmail.com abstract. this paper entails an empirical and computational assessment of the air flow field in the close proximity of a vertically positioned radiant cooling panel. this radiant cooling solution differs from the conventional large-area radiant cooling systems (e.g., ceiling panels). it involves rather small-sized vertical panels positioned close to occupants. moreover, the panels are designed so as to manage potential surface condensation of water vapor via integrated drainage elements. hence, the panels can be operated with relatively low surface temperatures. the low panel surface temperature and its proximity to the occupants are intended to compensate for the potential lower cooling power due to the relatively small panel size. in this paper, we specifically explore the air flow field close to the local radiant cooling panel via laboratory measurements and cfd (computational fluid dynamics). thus, possible issues regarding discomfort due to draft and turbulence risk close to the radiant panel can be examined. to this end, a prototypical local radiant cooling panel was installed in a mock-up office room of a laboratory. during the experiments, the air flow speed was measured and simulated at several heights (between 10 and 110 cm from the floor) and distances (ranging from 1 to 50 cm from the radiant panel). the results allow for the evaluation of the draft discomfort risk as well as the reliability of cfd in reproduction of the measurement results. a further step involved the numeric analysis of the effect of the human model on the air flow patter. keywords: radiant cooling panel, thermal comfort, air flow, cfd simulation. 1. introduction and background it has been suggested that climate change due to greenhouse gas (ghg) emissions is one of the most significant challenges facing humanity [1]. specifically, the global reliance on non-renewable energy sources leads to an increasing concentration of greenhouse gases in the atmosphere and thus contributes to climate change. in this respect, particular attention should be given to the household sector as a major consumer of energy [2]. the move from uniformly conditioned indoor environments to smaller, individually controlled zones, can be advantageous in view of thermal comfort and energy efficiency [3]. limitations of traditional cooling systems, such as noise, draft, vertical thermal gradients, and high energy demand necessitate the investigation of alternative cooling systems. in this context, the application of water in chilled radiant panels can be of interest [4]. these panels can be broadly classified based on their installation location. the most common form is the ceiling-mounted cooling panel. radiant cooling ceiling systems were investigated, among others, in the laboratories in european countries in the early 1990s [5]. they have been shown to have the potential to provide adequate thermal conditions [6]. in some systems, the pipes are inserted in the surface layer of a wall, ceiling, or floor [7]. lately, the integration of radiant cooling into furniture (e.g., in desks) has received attention [8]. thereby, the chilled water could circulate through the table top to remove the sensible heat, while latent heat is removed by an all-air cooling system [6]. as compared to conventional cooling systems, radiant cooling panels have the potential to offer smaller vertical temperature gradients, lower air movement velocity, and reduced local discomfort [9]. whereas conventional air-conditioning systems are mainly based on convection, radiant panels utilize a combination of radiation and convection [10]. the thermal performance of the cooling panels has been investigated via both experimental and computational methods. the thermal comfort of five human subjects in a laboratory test room equipped with cooling ceilings was analysed by nagano and mochida [11]. in their experiment, the temperature difference between the ceiling and the room air was less than 5 k. according to their findings, the mean radiant temperature for a supine human body should be applied in the design of ceiling radiant cooling. another study compared the radiant cooling systems with airconditioning systems in terms of thermal comfort, energy consumption, and cost. the results suggested that the radiant ceiling panel system is capable of creating a smaller vertical variation of air temperature and a more comfortable environment [12]. experimental studies are of course highly instructive, but also rather expensive and time-consuming. if carefully applied, cfd-based numeric simulation can provide alternative or at least complementary means of investigation. computational fluid dynamics has 269 https://doi.org/10.14311/app.2022.38.0269 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en s. najaf khosravi, h. teufl, a. mahdavi acta polytechnica ctu proceedings become an important tool in the prediction of thermal comfort in occupied spaces [13]. compared to the experimental techniques, cfd simulations can provide detailed values regarding the flow distribution and concentration fields in the whole domain, rather than just targeted points for data collection [14–16]. nevertheless, proper cfd application necessitates sufficient grid resolution, accurate choice of numerical models, estimation of numerical errors, and control over other numerical parameters for precise verification and validation [17–19]. despite the dramatic progress in its application, cfd has not replaced experimental and theoretical analyses, but represents a highly useful complementary tool [20–22]. as such, experimental studies have been conducted to probe the precision of numerical simulation of the indoor thermal environment [23]. myhren et al. [24] applied cfd simulations for two office rooms equipped with separate heating and ventilation systems, to investigate potential cold draught problems, the differences in vertical temperature gradients, and air speed levels. another study compared the energy efficiency of the cooling panel with the all-air cooling system via cfd simulation and experimental method. thereby, the cooling panel was found to be more energy efficient [25]. predicting thermal comfort by applying virtual human manikins in the cfd simulation has been reported by numerous studies [13]. according to our observation, there are no relevant standards in cfd simulations for the sizes, shapes, and postures of human model, and their application mostly depends on the purposes of the research works [26–29]. because of the excessive simulation time, the majority of cfd studies have applied simple (e.g., rectangular) shapes. in some cases, the nature of the study necessitated a complex human model [30]. within the framework of a recently completed research study, the potential of vertical radiant panels in the proximity of users was investigated and related technological requirements regarding envelope tightness and ventilation systems as well as water vapor condensation risk were evaluated [27]. as mentioned before, the personal radiant cooling panels are designed so as to manage potential surface condensation of water vapor via specific elements. one option is to remove the condensed water via a drainage system. however, in our experience, a much simpler option would be sufficient in most case, namely the placement of an easy to clean metal container underneath the vertical cooling panel. in the present contribution, we explore the potential of cfd simulations to analyse the airflow patterns and the velocity fields inside a test room with a vertical radiant panel positioned in the close proximity of a workstation. simulated results were compared with previously obtained measurements. the objective of the numeric analysis was to capture the air flow patterns and local discomfort risk. the study included also the air velocity field around a simplified human model. figure 1. test room equipped with a cooling panel. 2. approach predicting the thermal performance of a cooling panel is not a trivial task. in general, this system represents a mixed convection heat-transfer mode, including human body buoyancy-driven thermal plumes, and radiation through the panel and wall surface. to this end, a prototypical local radiant cooling panel was installed in a mock-up office room of a laboratory (see figure 1). to deal with potential surface condensation, a metal container was positioned below the local cooling elements. during the laboratory experiments, the ambient air temperature and relative humidity in the room were kept at 30 ± 0.5 °c and 40 ± 3 %, respectively. the target panel surface temperature was 10 °c. for these indoor air and panel surface temperatures, the cooling power of the radiant element was estimated to be roughly 100 w per square meter panel area [27]. the air flow speed was measured and simulated at several heights (10, 35, 60, 85, and 110 cm above the floor) and distances (1, 1.5, 2, 3, 4, 5, 10, 15, 20, 30, 40, and 50 cm from the radiant panel) (see figure 2) [31]. in this study, the potential discomfort risk due to draft was explored and was not found to be significant [31]: at a distance of 10 cm and more from the panel, the air flow velocity was never found to be above 10 cm s−1. the cfd model’s boundary conditions were adapted from the thermal conditions of the laboratory. note that during the experiment the surface temperature of the panel’s frame was higher than the chilled part of the panel. within the cfd model, the cooling element was simplified and no distinctions were made between the surface temperature of the frame and the chilled part of the panel. as a result, for the cfd simulations the whole panel temperature 270 vol. 38/2022 measurement and cfd analysis of a local radiant cooling solution figure 2. floor plan and sensors locations at the test room. was considered to be 11 °c. cfd model was implemented in the finite volume code ansys fluent 19.0 [32]. the model involves the vertical cooling panel mounted in the test space and a simplified human model for some simulation cases. ansys design modeler and ansys meshing were applied as a pre-processor to create the geometry, mesh, and the computational domain. in this model, a mesh with hexahedral element was generated. buoyancy due to density change was applied in the energy equation. the discrete ordinates (do) model was applied as the radiation model. the pressure-based solver was applied. the simple segregated solver was used for pressure-velocity coupling. various studies have assessed the performance of turbulence models. for rans-based models, selecting a proper turbulence model is essential for heat transfer analysis. the answers to the question of the best-performing turbulence models are not always consistent. for instance, the standard k-ε or the rng k-ε turbulence models are widely implemented for cfd simulations of cooling panel, while sst-k ω model has been applied in other studies. in addition, some studies applied a laminar flow model. the results of the literature review revealed that the k-ε family of turbulence models presented similar results consistent with laboratory data [33–35]. the skw and sst turbulence models showed a lesser performance level in predicting air velocity [36, 37]. this disparity can be due to the physical features of the case studies, as well as different computational parameters and settings applied in different studies. hence, for each specific simulation study, the selection of the turbulence model may have to be reassessed. in the present study, cfd simulations were carried out using the same geometry but with different turbulence models (rng k-ε and sst-k ω). to check the reliability of the turbulence model, the simulation results were compared with experimental. note that we used the discretization scheme employed for the energy, momentum, turbulent kinetic energy, and specific dissipation rate was the second order upwind scheme. pressure staggering option (presto) was used for the pressure discretization scheme. 3. results and discussion the air velocity could be an influential parameter for the thermal comfort: increased airspeed can aid the evaporation of sweat thus leading to a cooling effect, particularly if loose clothing is worn. nevertheless, too high air velocity may cause discomfort and a draughtiness sensation. average air speed is typically recommended not to exceed 0.15 m s−1 in the occupied zone to prevent discomfort [38]. the cfd method can complement the experimental results by allowing the detailed examination of the velocity field at different spots. given the complexity of the physical processes involved, a candidate simulation model should ideally incorporate complete three-dimensional processes, suitable turbulence models, realistic boundary conditions, and temperature-dependent material properties. in this study, to gain confidence concerning the model’s reliability and the employed turbulence model, we first compared the computational results with the laboratory measurements. subsequently, the study covered the human model analysis. 271 s. najaf khosravi, h. teufl, a. mahdavi acta polytechnica ctu proceedings figure 3. velocity as a function of distance from the panel for two different turbulence models. figure 4. simulated velocities plotted against corresponding measurements for two different turbulence models. 3.1. turbulence model two different turbulence models (k-ε and k-ω) were applied for the cfd simulation. during the experiments, air velocity was measured at 75 locations. figures 3 and 4 illustrates the velocity distribution profile (on the sensors location) obtained from the cfd simulation for two different turbulence models, together with the measured velocity values in the test room. the effect of turbulence models on the air velocity distribution can be seen in this figure. as such, the k-ε and the k-ω family of turbulence models overestimated the values by up to 0.05 and 0.1 m s−1 respectively. in this study, the cfd simulation results obtained while using the k-ε family of turbulence models were found to more closely correspond to the experimental results. the air velocity values were below the limits of the standards at the assumed seating position of the occupants [38]. as mentioned earlier, in this study we considered a highly simplified human model (body temperature 35) in three distances from the cooling panel (20 cm, 35 cm and 55 cm). the cooling was provided by radiant panels and the manikin acted as the source of heat generation in the room. the flow in the room is dominated by buoyancy. the solver settings were identical to those used for the aforementioned comparison with the experimental results. the main objective of this query was to explore the potential influence of the human presence on the air flow pattern around the radiant cooling panel. figure 5 illustrates the simulated velocity contours for the aforementioned simulation scenarios. the results indicate that the velocity in the vicinity of manikin lies in the range of 0.10 to 0.28 m s−1. the velocity range in the zone can be suggested to be 272 vol. 38/2022 measurement and cfd analysis of a local radiant cooling solution figure 5. simulated velocity contours around a human model at three distances from the radiant panel, namely distance of human model from panel, namely 20 cm (a), 35 cm (b), and 55 cm (c). within the recommended maximum velocity range to avoid draft [39]. with increasing distance from the cooling panel, its effect on the velocity field clearly diminishes. 4. conclusion using local radiant cooling panels has been suggested to be a potentially promising alternative to 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[39] l. berglund, a. p. r. fobelets. subjective human response to low-level air currents and asymmetric radiation. ashrae transactions 93(1):497–523, 1987. 274 https://doi.org/10.1016/j.buildenv.2006.04.027 https://doi.org/10.1016/j.buildenv.2009.06.008 https://doi.org/10.1016/s0360-1323(01)00029-4 https://doi.org/10.1016/j.ijheatmasstransfer.2013.02.005 https://doi.org/10.1016/j.ijheatmasstransfer.2013.02.005 https://doi.org/10.1146/annurev.fluid.29.1.123 https://doi.org/10.1016/s0376-0421(02)00005-2 https://doi.org/10.1016/j.buildenv.2005.06.029 https://doi.org/10.1111/j.1600-0668.2011.00723.x https://doi.org/10.1111/j.1600-0668.2011.00723.x https://doi.org/10.1007/s12273-015-0212-0 https://doi.org/10.1016/j.buildenv.2015.02.031 https://doi.org/10.1016/j.enbuild.2007.04.011 https://doi.org/10.1016/j.buildenv.2004.04.010 https://doi.org/10.1016/j.buildenv.2004.01.026 https://doi.org/10.1016/j.enbuild.2021.111104 https://doi.org/10.1080/10789669.2005.10391156 https://doi.org/10.1016/j.buildenv.2017.03.040 https://doi.org/10.1177/1420326x13480056 https://doi.org/10.1088/1742-6596/2069/1/012122 https://doi.org/10.1007/bf01061452 https://doi.org/10.1016/0045-7930(94)00032-t https://doi.org/10.1016/0017-9310(72)90076-2 https://doi.org/10.2514/3.12149 https://doi.org/10.1016/s0378-7788(98)00020-6 acta polytechnica ctu proceedings 38:269–274, 2022 1 introduction and background 2 approach 3 results and discussion 3.1 turbulence model 4 conclusion references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0539 acta polytechnica ctu proceedings 38:539–545, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the potential of unused railway areas jiří kugl czech technical university in prague, faculty of civil engineering, department of urban design, town and regional planning, thákurova 7, 166 29 prague 6, czech republic correspondence: jiri.kugl@fsv.cvut.cz abstract. railway transportation faces many of the issues that are related to standard brownfields – due to changing technologies and industries, more effective systems (of traffic control in this case) and evolving needs many of the areas are actually no longer necessary for proper function of the railway. that is especially prominent in the case of railway stations, where a significant number of the stations use just a portion of their available tracks, buildings and areas (for example for unloading or repair). the remaining areas are sometimes sporadically used, but more effective and conscious management of the station could fairly easily lead up to the release of these areas for another, more useful function for both the city and the station. this paper will explore the differences between “standard” brownfields and unused railway areas like the typical ownership structure, particular location within the city and the effect station has on the city structure, composition and topography of the areas or the fact that railway areas are most often never really fully abandoned and they do continue to serve in some, albeit diminished capacity. paper also aims to map out how much of the railways areas are underused or unused in czech republic (country with highest rail network density in the world). this paper will then recommend the best ways to use and revitalize them and it will show some successful examples of revitalization projects from all around the world. keywords: railway, unused, brownfield, revitalization. 1. potential of unused areas finding solutions of the problem of unused areas, resp. brownfields have long been recognized worldwide as essential for the healthy sustainable development of cities. the use and revitalization of brownfields is one of the answers to the fundamental problem of undesirable and unsustainable growth of cities in the suburbs and the related occupation of land. there is a shift in land perception, when it is now generally considered to be one of the finite resources with all the consequences that follow (there are also opinions that the land is actually one of the most valuable finite resources, as it is a source of livelihood not only for humans, but also for other organisms). the revitalization of unused areas is closely related to current issues of urbanism and spatial planning, namely suburbanization and the so-called urban sprawl (uneconomical, often unconceptual and parasitic development in the suburban landscape). the development of unused railway areas has the potential to dramatically reduce this phenomenon and in addition, it has many advantages over building on the so-called “greenfield” – such as the use of above-standard connections to existing transport and technical infrastructure of the city, healing “wounds” in the urban fabric, suppression of the barrier effect of railway, remediation of contaminated soils and thus the improvement of the environment, etc. the development of unused areas also addresses their negative impact on the surrounding area, whether in the form of a direct impact of architectural and urban degradation of the environment or barrier effect, or in the form of an indirect effect such as the fall of property prices, an overall outflow of funds from the territory potentially leading to an increase in unemployment, an increase in vandalism and crime or a degradation of the aesthetic taste of the population. for not completely abandoned, only not fully utilized railway areas, these indirect consequences are usually less noticeable than for traditional brownfields, but they are certainly not negligible. as a result, the development of unused land has a double basic justification – it not only helps to solve the problem of new development, but also reduces the negative impact of the unused land itself. here it is necessary to emphasize the lesson from practice that many negative processes can be reversed – for example, a gradual outflow of funds from a neglected area, where a quality pilot revitalization project can (and often does) serve as a catalyst for a significant change in the whole area (the best-known example of this approach in the field of railways is the famous high line park in new york). 2. specifics of the unused railway areas first, it is necessary to define what we consider to be unused railway areas and what separates them from the usual term – brownfields. brownfields are defined by czechinvest as: “a property (land plot, building, complex) that is disused and neglected and may even be contaminated. brownfields are remnants of industrial, agricultural, 539 https://doi.org/10.14311/app.2022.38.0539 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en jiří kugl acta polytechnica ctu proceedings residential, military or other activities. a brownfield cannot be used effectively and appropriately without undergoing a process of regeneration.” [1]. calling unused railway areas brownfields would therefore be potentially problematically misleading, as it gives a wrong impression of a narrower focus only on completely disused and abandoned railway areas (because this is how brownfields are generally perceived, at least in the minds of the general public). however, it should be noted that the issue of revitalization of unused railway areas of course shares much in common with the issue of brownfield revitalization. for the purpose of this research the use of the term “unused railway areas” is proposed which gives a needed clearer idea of the character of these areas with no ambiguity. unused railway areas are then defined as follows: “these are railway areas that haven’t been in use for long time or that are used very inefficiently (i.e. although they are sporadically used or railway operates there still to some extent, by more rigorous application of the principles of modern railway operation these areas could easily be released without detriment to the railway function).” compared to standard brownfields, unused railway areas also differ in the following specifics: the role of the railway railway played an important role in the development of the cities through which it passed. it affected both their industry (by connecting it to national and international railway network and thus facilitating the transport of materials and resulting products), as well as the economic situation of the city itself (overwhelmingly positively) and its urban structure. the specific location of the railway and especially the station had a major impact on the direction and manner of subsequent development of the urban structure. the track as a linear (and in the case of larger mainly freight stations also area) barrier often determined the development zones, but also affected the functional structure of the city. railway and significant expansion of industrial areas was often connected, when in some cases special sidings (or even their network) were introduced into individual factories. in many cases, railways and railway stations were built at a considerable distance from the existing city for technical and economic reasons, however, this mostly became a major impulse for the development of the city towards the station, which was soon absorbed by the urban structure. the result was often a de facto shift in the position of the station from the suburbs to the inner parts of the city, when the station sometimes reached the central positions of the newly formed districts. the arrival of the railway to the city was also one of the impulses leading to the demolition of the medieval fortifications. the stations at this time functioned as new symbolic (but also very much actual) gates, entrances to the city. they became the new starting points of the main flow of visitors and passengers in the city. the builders of the station buildings were usually aware of their importance (especially in the golden era of network construction in the second half of the 19th century) and reflected it with a representative architectural design of the station buildings. the structure of the city reacted to this in many respects in a similar way, as if it were a city gate. the protection zone of the track according to the law also plays an important role. the railway is also a significant element in landscape, whether in the form of the track itself or civil engineering works and modifications such as notches, embankments, retaining walls, tunnels, galleries or bridges. and again, it should be mentioned that it effectively forms a linear barrier in the landscape. this effect is more noticeable (to some extent also spatially, but above all functionally) the stronger and faster the operation on a particular track is. single-track local railways thus form a somewhat delicate barrier and are relatively easy to cross (by people and animals), on the other hand, the corridors of interstate trains (not to mention high-speed lines) are practically insurmountable and thus significantly divide the landscape. similarly, the aesthetic effect of the railway in the landscape varies, where some elements such as viaducts can generally be considered as an interesting and distinctive local feature, however, a large freight station is not as a positive element in the landscape from any point of view. great transport infrastructure connections most of these areas, by the logic of their function, have a very good connection to public transport, sufficient parking areas and are very well connected to the road infrastructure. “strategic” interests of the state in some cases, the state may be interested in maintaining the selected line even though the operation has not been carried out on it for a long time and the line is pretty much abandoned. reasons are usually to ensure the state’s defenses. the nature of the area considering the size of the areas, there is a minimum of buildings and objects in general, most of the area consists of tracks. dimensional character usually longitudinal or strongly longitudinal shape. thanks to obvious technical parameters, railway forms an uninterrupted linear barrier. ownership structure with a few exceptions, there is one, maximum two owners, which greatly simplifies the revitalization process. in the czech republic, for example, there are typically two owners, the state-owned české dráhy (especially some railway stations and the land below them) and the national správa železnic (sž) (everything else) – considering those are state companies, it makes the possible different future use of these areas by the cities easier 540 vol. 38/2022 the potential of unused railway areas and more feasible. after the public sale of part of the railway properties, some unused railway areas fell into the hands of private owners (mostly small family houses for railway personal, warehouses and land in the vicinity of the railway lines). an exception is sidings in industrial areas, which have private (and therefore often very confusing) ownership. railway station architecture the station buildings were very often, by their nature, built with an emphasis on high representativeness. in the cases of the most important stations, original architecture was usually applied (to this day with a very high aesthetic value), at other stations, standardized buildings were designed and built. contamination railway operation causes some typical organic and inorganic contamination of the site. the sources are mainly waste from lubricating oils and condenser fluids, the transport of petroleum derivatives, metal ores, fertilizers and various chemicals or the use of herbicides [2]. the most frequently detected contaminants along railway lines are metals, pesticides and components of crude oil or fuels (petroleum products). these chemicals are associated with standard track operation and are likely to be found anywhere along the track [3]. the three most important types of contaminants are polyaromatic hydrocarbons (pahs), heavy metals and polychlorinated biphenyls (pcbs) (to some extent) [4]. terrain the railway is usually located on a very flat terrain without sudden breaks, there is a minimum of vegetation in the track area. 3. extent of unused railway areas in the czech republic research of the railways in the country was carried out which aimed to identify how many unused railway areas are located in the czech republic and what revitalization potential is hidden in these areas. the analyzed sample is based on the sž document “program of reconstruction and revitalization of passenger stations for the years 2018-2022” [5], which was then manually supplemented with some important stations, which were not originally included in the document (because no investment was planned in those buildings in 2018-2022) and several larger freight stations. the final sample was then 712 stations (604 from the original list of sž, 108 added). in the czech republic in 2021 there are a total of 2 612 railway stations and stops. there are 1 082 railway stations and 1 530 stops (typically only a shelter) [6]. the analyzed sample of railway stations therefore makes up about 66 percent of all railway stations in the czech republic. the sample is therefore not exhaustive, however, due to the fact that most of the remaining stations are mainly stations in smaller and very small cities and villages, it is clearly sufficiently representative to identify the overall extent and potential of unused railway areas in the czech republic. the analyzed parameters in the research includes, for example, the position of railway areas within the city, the size of the city, the size of railway areas (number of buildings, tracks, sidings) and especially the revitalization potential of these areas. it is based on the categorization according to the economic profit analysis from the usa that is used today in the czech republic for categorizing brownfields [7]. this categorization was slightly adapted to focus on railway areas and their revitalization potential. the evaluation is qualitative, it is an expert evaluation of many factors, such as the position of the railway area in the city, its connection to other forms of transport, the size of the city itself (and thus its economic strength), composition and location of existing functions in the city and its civic amenities (looking for possible deficit that could be solved by the revitalization of the railway area), the size of the unused railway area (number of the buildings, number of tracks or empty part of the land formerly used for unloading material), its character (building or area) and its location within the station, architectural the quality of the buildings, what the rest of the station looks like and how it is used, how intensive traffic on the track takes place (according to the number of passengers), what the neighboring buildings look like and what is their function or where the development of the city can be expected (because the revitalized railway areas could for example play an important role as a central linking element there). an important monitored factor is also the existence of other unused areas (brownfields) within the residence, such as not fully used castle, representative factory buildings or large farms (in a smaller cities or villages), in short buildings where higher priority of possible revitalization projects can be expected (whether from urban, cultural or aesthetic point of view). it is necessary to mention that some things are not part of the research although they do matter – for example, how many spaces in the station are successfully rented in the long run or where, for example, the station staff lives. those information are too detailed for the purpose of this research which was to get a general idea of the extent of unused railway areas in the country, not to evaluate any particular railway are in great detail. as the figure 1 shows, the research found that in the czech republic, 74 percent of the analyzed stations are located in cities of less than 10 000 inhabitants. the majority (57 percent) of stations are located in central positions within the city and as much as 65 percent, or about two thirds of all stations are located in the inner parts of the cities. this proves the assumption that large number of railway areas do have a very important and prominent position within the city. only 12 percent of the stations are located in an external position, in an area outside of the city. 541 jiří kugl acta polytechnica ctu proceedings figure 1. distributions of railway areas (author’s research). revitalization potential could be identified in as many as 40 percent of stations (288 stations), 87 stations have a high potential a (revitalization of these areas is appropriate and recommended) and 201 stations have a potential b (revitalization of these areas is possible). it is interesting to note that the highest number of stations with a high potential for revitalization (rating a) is located in the smallest cities with a size of less than 1 000 inhabitants. that is partially caused by the uneven character of the analyzed sample and the railway station network in czech republic, where sizable portion (25 percent) of stations is in towns under 1 000 inhabitants. other reason is the fact that the revitalization potential evaluation scales with the size of the city – for example transforming part of the unused train station into a local cafe with small gallery plays a significant role in the life of a small town, but doing the same project in a big city would have negligent effect and therefore the investment would not be warranted. another important thing to note is the fact that for cities larger than 10 001 inhabitants, there is usually a 60–72 percent probability that their railway areas will have a revitalization potential. in comparison, in cities with a size of less than 1 000 inhabitants the probability of a revitalization potential is only around 34 percent. unused railway areas with rating a (great revitalization potential) have an average of about 13 tracks and a median of 9, an average of 10 sidings and a median of 5, an average of 9 buildings, a median of 4, therefore they are rather large stations. rating b railway areas then have an average of about 11 tracks and a median of 7, 7 sidings and a median of 3, an average of 7 buildings, a median of 4, which are medium to large stations. a typical station in the analyzed sample has an average of about 8 tracks and a median of 5, an average of 5 sidings and a median of 3, an average of 5 buildings, a median of 3. this information indicates that stations with revitalization potential are usually large and therefore highly likely to have large unused areas within them, which represents a significant revitalization capital for individual cities. unsurprisingly, the highest number of stations with a high potential for revitalization is in central positions within the cities or at least their districts. by correlation calculations (using pearson correlation coefficient) of the different parameters of the research, it turned out that the most significant factor in evaluating the revitalization potential was the position of the unused railway area within the city (which seems rather logical) and the number of buildings. an important finding is although the fact that no parameter has proven to be a major factor that would dramatically affect the evaluation of revitalization potential. the evaluation thus proved to be really multifactor and it is therefore necessary to monitor all of the set parameters. the research proved that there is a large number of unused railway areas in the czech republic with revitalization potential. these areas – taking into account their characteristics within the cities – represent a huge potential for improving the quality of life of local residents (e.g. as additions to civic amenities or as new public spaces) as well as a large treasury of areas and buildings which are owned by the state and are not used to their full potential. 542 vol. 38/2022 the potential of unused railway areas 4. possibilities of further use for any city, a quality revitalization of unused railway areas is absolutely essential, because as the research analysis has shown, these areas often play an important role in its entire structure. it is highly desirable to find solutions that allow these areas to be used meaningfully and re-integrate them into the fabric and life of the city. in general, it is also appropriate to reduce the barrier effect of the railway as much as possible (by reducing the number of tracks, building footbridges, underpasses, etc.) and try to maximize the permeability of the area (freeing up unused tracks, making them a public space preferably). also, the current assumption in the field of architecture and urbanism also applies that it is always better to use already built-up areas and existing buildings than to build new ones on open green areas (whether in terms of urban context, architectural-historical continuity or in terms of sustainability). it must be emphasized that the physical destruction of railway areas should always be preceded by careful consideration, as such action is against the european union’s transport policy, in particular the marco polo ii program (concerning the transfer of freight from congested roads to railways or waterways). however, many railway areas are decidedly unused and cannot realistically be expected to be re-used efficiently, in those cases it makes great sense to consider another use that will be more beneficial for the city (and, after all, also for the station itself). it is understandable that the size of the city and its budget (or the financial possibilities of sž or other railway carriers) significantly affects the possibilities of revitalization and thus the choice of further use of freed unused railway areas. however, it is still necessary to keep in mind that the use of these areas is in most cases the most financially viable way to supplement or enrich the network of civic amenities, as it is state property after all, where it can be expected (with the support of respective state authorities) there could be a gratuitous or at least very advantageous transfer of real estate into the ownership of the city. one of the possible ways of utilization of unused railway property is also its active lease (targeted at desirable functions), either by the city or sž (in coordination with the city and its vision). this applies, of course, mainly to buildings. an interesting option is also in justified cases the waiving of the lease for a definite period of time (e.g. for state or non-profit organizations or interest groups). in this way, it is possible to support, for example, clubhouses serving local communities, fitness and health groups, groups for children, necessary missing services such as barbershop, surgery, library, but also commercial activities such as start-up offices or social programs such as social apartments for young families or flats for the necessary professions (typically teachers, doctors). this can effectively serve to rather cost-effectively kickstart local community life. many possibilities for further use can also be quite minimalistic – it is understandable that a 3 000 city will not have enough funds to renovate the building into a modern, elegant hamburger bahnhof-style gallery, but sometimes a somewhat simple and industrial, yet functioning gallery (or cultural hall etc.) can be created by just cleaning the building, possible wiring or plumbing done very cost efficiently visibly along the walls and repair of the roof (if necessary). similarly, other creative ways of saving can be used – instead of reconstruction (for example of the former storage hall), the building can ideally drop part of one or more walls (at the very least the parts underneath the windows) and create an unusual generous outdoor shelter which can be used, for example, for outdoor seating or as a roof for holding markets. the costs of such projects are, of course, fundamentally different from a complete reconstruction with repair of the facade, all surfaces, replacement of windows and roofs, and many examples from around the world but also in the czech republic show that such areas have their specific charm, work very well and they usually find their users. from the point of view of costs (and, after all, the time of implementation), this minimalist approach can be applied to practically all possibilities of further use, including the creation of new public spaces. the areas can be done attractive, for example, by adding (as a kind of technical artifact) a parked wagon, which with minor modifications can easily serve as an interesting gazebo, or a lockable storage for a refreshment stand. former loading area doesn’t have to be repaved, it could be just creatively painted over using colors for marking roads, creating visually striking fun patterns, labyrinths or improvised “sport fields”. interesting and attractive in the case of revitalization of unused railway areas does not always mean costly, it could be inspired by the ideas of “tactical urbanism”, have a clear vision and consistent emphasis on low cost and finding ways to achieve it (e.g. looking for possible cooperation with local schools or sponsoring local companies). in general, practically all stations in the czech republic should ideally expand their range of services and shops (for which it is sensible to use those areas and buildings within the station that are not currently used effectively) and follow the global trend aimed at improving the quality of travel for passengers. one has to remember, that even today, for a large number of people, the station is still the gateway to the city, first point of entrance, and as such should be, if not (for financial reasons) representative, then at least lively and interesting. one things that should be noted is that the station buildings were very often, by their nature, built with an emphasis on high representativeness. the station architecture is quite specific and even most laymen are able to identify the station purely by its appearance. for practically all stations it is true that to this day 543 jiří kugl acta polytechnica ctu proceedings figure 2. possible new functions of unused railway areas [8]. they stand with great dignity within their cities and are often (especially in smaller cities) one of the most representative buildings in the area. the vast majority of stations were established during the golden age of the railway and therefore an unusual amount of funds was invested in them. by finding new use for these old buildings, the city can relatively cheaply obtain aesthetical and expensive buildings, far more representative than the city would usually have the funds for. many of the unused railway areas also pose a significant cultural, architectural or historic value. one of the possibilities is the determination of the building or area (or its part) a cultural monument and thus the provision of monument protection. the selection of railway monuments in czech republic was initially somewhat random and was accompanied by many disputes, when the declaration of a monument was cancelled. now, with the growing public interest in industrial heritage, the situation is slowly improving. in 2021, there are 250 railway areas on the central list of cultural monuments, of which about 124 are stations (out of 1 082 functional stations) [9]. when determining the future function of the unused railway areas, one of the major specifics of these areas has to be taken into account – in vast majority of the station, the operation is still running to some extent. revitalization projects therefore should work with this factor and use it to their advantage. the worldwide trend is towards the greatest possible involvement of the station in the daily life of the city, where many of the added or implemented functions improve the quality of life and services of both the neighbourhood and the railway operation itself. for unused railway areas, the following possibilities for further use should be considered (and in some cases they are mutually combinable or consecutive). the examples shown in figure 2 come from countries like usa, germany, france, australia and czech republic. 5. conclusion the research found that in the czech republic, the majority (57 percent) of stations from the analyzed sample (of 712 stations) are located in central positions within the city and about two thirds of all stations are located in the inner parts of the cities. thus proving the assumption that railway areas typically have a very important and prominent position within the city. only 12 percent of the stations are located in an external position, in an area outside of the city. in 40 percent of stations (288 stations) revitalization potential could be identified. from those 87 stations have a high potential a (revitalization of these areas is appropriate and recommended) and 201 stations have a potential b (revitalization of these areas is possible). size of a city of course plays a big role in evaluating the revitalization potential of the railway area, where for example in cities larger than 10 001 544 vol. 38/2022 the potential of unused railway areas inhabitants, there is a 60–72 percent probability of a revitalization potential. successful revitalizations of unused railway areas have the potential to have a very positive effect on their immediate surroundings and so, in addition to their obvious direct benefit (in the form of, for example, the introduction of a new function to the area), they also have an equally important indirect benefit. it is a practice many times proven that successful attractive projects can start a positive development of the whole area. however, even smaller revitalization projects have a significance that should be utilized. due to the fact that the vast majority of unused railway areas are state property (whether čd or sž), the use of these areas is in many cases the most financially viable way, for example, to supplement or enrich the network of civic amenities or public spaces. unused railway areas can be re-integrated into the life of cites, in which they can again play a significant and positive role, which the railways historically always did. references [1] czechinvest. brownfields. [2022-02-08]. https://www.czechinvest.org/en/forinvestors/real-estate-offer/brownfields [2] b. wiłkomirski. railway tracks – habitat conditions, contamination, floristic settlement – a review. environment and natural resources research 2(1):86– 95, 2012. https://doi.org/10.5539/enrr.v2n1p86 [3] department of environmental protection. best management practices for controlling exposure to soil during the development of rail trails. [2016-02-20]. https://www.mass.gov/doc/best-managementpractices-for-controlling-exposure-to-soilduring-the-development-of-rail-0/download [4] united states environmental protection agency. technical approaches to characterizing and cleaning up brownfields sites: railroad yards. [2016-02-20]. https://nepis.epa.gov/exe/zypurl.cgi?dockey= 30004h7u.txt [5] sždc. program rekonstrukce a revitalizace osobních nádraží pro roky 2018-2022, sždc, praha, 2017. [6] správa železnic. základní charakteristika železniční sítě. [2021-06-14]. https://www.spravazeleznic.cz/onas/vse-o-sprave-zeleznic/zeleznicecr/zeleznicni-sit-v-cr [7] b. j. jackson. příručka pro vlastníky brownfieldů, ústí nad labem, 2011. [8] j. kugl. revitalizace železničních ploch. in přestavba a regenerace sídel a krajiny, pp. 34–37. ústav územního rozvoje, brno, 2017. [9] národní památkový ústav. vyhledávání, ústřední seznam kulturních památek. [2021-06-15]. https://pamatkovykatalog.cz/ 545 https://www.czechinvest.org/en/for-investors/real-estate-offer/brownfields https://www.czechinvest.org/en/for-investors/real-estate-offer/brownfields https://doi.org/10.5539/enrr.v2n1p86 https://www.mass.gov/doc/best-management-practices-for-controlling-exposure-to-soil-during-the-development-of-rail-0/download https://www.mass.gov/doc/best-management-practices-for-controlling-exposure-to-soil-during-the-development-of-rail-0/download https://www.mass.gov/doc/best-management-practices-for-controlling-exposure-to-soil-during-the-development-of-rail-0/download https://nepis.epa.gov/exe/zypurl.cgi?dockey=30004h7u.txt https://nepis.epa.gov/exe/zypurl.cgi?dockey=30004h7u.txt https://www.spravazeleznic.cz/o-nas/vse-o-sprave-zeleznic/zeleznice-cr/zeleznicni-sit-v-cr https://www.spravazeleznic.cz/o-nas/vse-o-sprave-zeleznic/zeleznice-cr/zeleznicni-sit-v-cr https://www.spravazeleznic.cz/o-nas/vse-o-sprave-zeleznic/zeleznice-cr/zeleznicni-sit-v-cr https://pamatkovykatalog.cz/ acta polytechnica ctu proceedings 38:539–545, 2022 1 potential of unused areas 2 specifics of the unused railway areas 3 extent of unused railway areas in the czech republic 4 possibilities of further use 5 conclusion references https://doi.org/10.14311/app.2022.33.0098 acta polytechnica ctu proceedings 33:98–104, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague trc sandwich solution for energy retrofitting isabella giorgia colombo, matteo colombo, marco di prisco∗ politecnico di milano, department of civil and environmental engineering, p.za l. da vinci, 32 20133 milano, italy ∗ corresponding author: marco.diprisco@polimi.it abstract. concerning energy improvement of existing façades, a favourable system involves prefabricated multilayer panels, made of internal insulation core and outer textile reinforced concrete layers. it is a convincing alternative to external thermal insulation composite systems (etics) and ventilated façades, and it meets all the requirements for façade systems. the main advantage is the possibility to apply the panel using a crane, without any scaffolding. the paper considers two solutions: the former uses expanded polystyrene (eps) as insulating material; the latter substitutes eps with an innovative green insulation material made of inorganic diatomite. the paper aims at comparing the solutions in terms of mechanical properties of the components and behaviour of the composite sandwich at lab-scale level. numerical models, previously calibrated, will be instrumental for the discussion. keywords: energy performance, sandwich panels, textile reinforced concrete. 1. introduction in europe, building sector uses a large amount of energy (40% of the total) and produces relevant co2 emissions (36%). for this reason, the focus of many european policies is to improve energy efficiency [1]. each year, only approximately 1% of buildings are energy renovated, in the framework of a building stock that is old and energy inefficient, with the half of residential buildings dated before 1970, year in which the first european thermal standard was emanated [2]. hence, improvement in energy efficiency of buildings is a crucial topic. in multi-storey buildings, façades cover a large amount of the total outer envelope surface. hence, solutions devoted to the energy retrofitting of existing façade result particularly effective in terms of building energy saving. a convenient solution is represented by multilayer precast sandwich panels made of an inner insulation core and thin external textile reinforced concrete (trc) layers. trc sandwich panels were proposed about 15 years ago by german researchers [3] and constituted a new idea in the framework of cement-based sandwich elements: trc thickness (10-15 mm) is smaller than that of traditional r/c external claddings (>40 mm) [4], leading an important decrease in the weight of the panel. other conveniences are durability and the good appearance, obtained using fine-grained mortar. although connectors were introduced between the layers for a durable connection [5], advantage is taken from adhesive bond between trc and the insulation layer to transfer the shear stresses. 2. trc sandwich solutions the research concerns prefabricated sandwich panels constituted by thin external trc layers and an internal insulation core. large size panels (3 × 1.5 m2) have been designed to reduce the manufacturing and installation costs. the installation is performed through a crane, without scaffolding, thus minimizing installation timing and inconveniences for inhabitants. other advantages are: the integration of the insulation and finishing systems; the durability and possibility to obtain different colour and texture as finishing; the protection of the insulation core through the external, waterproofed layer of trc; the good resistance to concentrated loads. only the inner trc layer is connected to the existing r/c frame; the shear transfer between the two trc layers is entrusted to the insulation material (for fire safety, some supplementary connectors could be inserted). 2.1. panel with expanded polystyrene core (easee project) the consortium of the european project easee [6] developed a prefabricated trc sandwich panel with expanded polystyrene (eps) as insulation material. a thickness of 100 mm was considered for the core. all the details concerning the panel design, testing at different scale levels, durability, and modelling can be found in [7–12]. easee main output is represented by a demonstration building [11] in cinisello balsamo, italy, where a 70% decrease in the thermal transmittance of the wall was achieved [13]. one of the main criticisms of this panel is that eps may melt during fire, leading to the loss of mechanical performances of the composite solution. 2.2. panel with diatomite core (smart p.i.qu.e.r.) smart p.i.qu.e.r. project [14] was aimed at solving the criticisms of the previous project. among the 98 https://doi.org/10.14311/app.2022.33.0098 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 trc sandwich solution for energy retrofitting figure 1. test results: trc tensile behaviour (a) and core material compressive behaviour (b). goals, particular relevance is assumed by the development of a green insulation material, with suitable mechanical properties and good fire and insulating capabilities. at the same time, trc layers need to be optimised, reducing the shrinkage and weight of the mortar, and enhancing the fabric fire behaviour and sustainability. 2.3. materials 2.3.1. textile reinforced concrete trc developed in the easee project is made by high strength fine grain mortar reinforced with one alkali-resistant glass fabric. the warp is aligned in longitudinal direction. this fabric [7] (leno-weave production; sbr coating: water resin based on styrene butadiene rubber) is characterised by a nominal tensile strength in warp direction of 820 mpa (computed on the equivalent cross-sectional area of the glass). the cementitious matrix has a water to binder ratio equal to 0.225, a superplasticiser to cement ratio of 9.3% and 1 mm maximum aggregate size. the average cubic compressive strength (fcc) of mortar, measured as specified in [15], is equal to 71.89 mpa on 12 specimens, with a coefficient of variation of 9.13%. concerning smart p.i.qu.e.r. project, several mortar mix-designs and alkali-resistant glass fabric reinforcements were investigated [14]. in this paper, a high-performance mortar that includes shrinkage reducing admixtures and is reinforced with one alkaliresistant glass fabric coated with a thermosetting epoxy coating is considered (see "hpc+sra_ap1" in [14]). the epoxy coating is particularly suitable in case of fire exposure. the fabric is characterised by a nominal strength in the warp direction of 1041 mpa. the cementitious matrix used is characterised by a maximum aggregate size of 1 mm, a water to binder ratio of 0.20 and a super-plasticiser to cement ratio of 5.1%. the average cubic compressive strength (fcc) of mortar measured according to [15] is equal to 97.53 mpa on 32 specimens, with a coefficient of variation of 8.33%. in both cases, three 400×70×10mm3 trc samples were tested in uniaxial tension using an electromechanical press instron 5867 with a maximum load capacity of 30 kn, imposing a constant crosshead displacement rate of 0.02 mm/s. the same test apparatus and modalities described in [7] are used. the test results are shown in figure 1a in terms of nominal stress (σ = p/a, with p=load and a=specimen cross section) vs. normalised displacement (δ/l, with δ = imposed displacement and l = free length) curves. 2.3.2. core material in the framework of the easee project, expanded polystyrene was used as insulation material in sandwich panels. the product eps250, characterised by a nominal density of 35 kg/m3, was selected. as resulting from experimental tests [10], it is characterised by a uniaxial compressive yield stress of 0.188 mpa, an estimated young’s modulus in compression of 13.7 mpa, a uniaxial tensile yield stress of 0.392 mpa, a shear yield stress of 0.160 mpa and an estimated shear modulus of 5.04 mpa. as already anticipated, one of the main aims of the smart p.i.qu.e.r. project was to develop a core material that could be more sustainable and more suitable in case of fire with respect to eps. hence, a novel lightweight sustainable insulation core (dhcf) was developed by verdolotti and other authors [16] using diatomite (natural source) as matrix. short fibres were added to guarantee a ductile behaviour of the insulation material both in compression and in bending, in order to guarantee a proper global behaviour of sandwich panels. these fibres also contribute to the enhancement of thermal performances of the material, by reducing the thermal conductivity. polypropylene (pp) fibres were selected among different kind of reinforcement tested (jute, waste synthetic, mixed and hemp fibres, as such and chopped). 99 i. g. colombo, m. colombo, m. di prisco acta polytechnica ctu proceedings figure 2. four-point bending test set-up (smart p.i.qu.e.r.). this choice was related not only to the proper thermal and mechanical performances of the foam in which pp fibres are included, but also because they are the only ones that allow the production at industrial level by means of an industrial plant. considering that the level of the plastic plateau is related to specimen density, a target density of 300 kg/m3 was fixed to guarantee comparable performances of the core material with those of expanded polystyrene. the core mix design is collected in table 1. diatomite (wt%) 19.57 polysilicate (wt%) 65.24 si (wt%) 0.046 catalyst (wt%) 8.39 vegetable (wt%) 5.82 pp fibres (wt%) 1 table 1. dhcf_pp mix design. compressive tests were performed on eps specimens with a size of 100 × 100 × 150 mm3, and dhcf specimens with a dimension of 40 × 40 × 40 mm3. the tests were displacement controlled by imposing a constant displacement rate of the machine crosshead equal to 1e-3 mm/s and 2e-3 mm/s respectively. the compressive behaviour of both expanded polystyrene and diatomite-based insulation material is shown in figure 1b where the nominal stress σ vs. nominal strain ε curves are plotted (σ = p/a, p = load, a = unloaded specimen cross-section; ε = δ/h, δ = top displacement, h = specimen height). 3. experimental behaviour of lab-scale sandwich beams with diatomite-based core a wide experimental campaign on lab-scale sandwich specimens were developed within the easee project. a four-point bending set-up was adopted for both deep (550 × 150 × 120 mm3) and slender (1200 × 300 × 120 mm3) sandwich beams, with eps 100 mm thick. detailed description of experimental procedures and results can be found in [8]. hence, this section is focused on the experimental tests performed at lab-scale level within smart p.i.qu.e.r. on beams in bending. 3.1. test set-up four-point bending tests were performed on sandwich beams using the same electromechanical press instron 5867 already mentioned. two nominally identical specimens with size of 600 × 150 × 80 mm3 and the warp of the glass fabric aligned in the beam longitudinal direction were tested. the specimen geometry and the test set-up are shown in figure 2. the insulation panel densities are collected in table 2. specimen ρ1st_measure ρ2nd_measure(kg/m3) (kg/m3) piquer1 330 (31 days) 303 (125 days) piquer2 348 (28 days) 297 (122 days) table 2. density of the insulation core. the insulation layer of full-scale panels designed in the smart p.i.qu.e.r. project is obtained placing side by side different core panels with size of 600×600×100mm3. for this reason, some cuts are introduced in the insulation layer also in the lab-scale beams (figure 2). with the aim of reducing stress concentration, 50 mm wide steel plates are glued on the specimen over the cylindrical supports and under the loading-knives. the loading and the supporting cylinders are free to rotate on their axis and in the plane perpendicularly to the longitudinal axis of the specimen. during the initial phase of each test, specimens were instrumented with displacement transducers in order to measure: the specimen deflection under the loading knives, the deflection of the upper trc layer under the loading knives, the superior longitudinal displacement on the compressed side inside the constant bending moment region, and the crack opening displacement on the tense side. however, these aspects are not treated in this paper. tests were displacement-controlled, considering the machine crosshead displacement (stroke) as feedback parameter. an initial stroke rate of 2e-3 mm/s 100 vol. 33/2022 trc sandwich solution for energy retrofitting figure 3. experimental results (smart p.i.qu.e.r.): load vs. stroke curves (a) and pictures of specimen piquer 2 during the test (b) and at failure (c). was imposed; then, after crack formation, the rate is increased up to 5e-3 mm/s. 3.2. experimental results the test results are shown in figure 3a in terms of load (p) vs. stroke curves. looking at the results, it is interesting to note that specimen piquer1 experienced a premature shear failure of the core, thus leading to a localised failure and preventing the multicracking of trc layers that guarantees the hardening behaviour of the sandwich solution. in fact, while performing a check before testing, core panel used to cast specimen piquer1 presented to the touch some softer regions at the edges. on the contrary, specimen piquer2 showed a ductile hardening global response, experiencing multi-cracking of trc faces. a picture of this specimen during testing is shown in figure 3b. the final failure is related to the debonding between the beam core and the lower trc layer (figure 3c). 4. numerical modelling of the tests and comparison between the two sandwich solutions 4.1. geometry and constraints abaqus /cae 6.14-5 was adopted for the modelling of the panel. yz symmetry was exploited, and half of the beam was reproduced, thus minimising the numerical effort (figure 4a). as shown in the figure, the panel layers (trc and core), and the steel plates are modelled as solid and homogeneous. perfect bond is assumed at core/trc interfaces and between the panel and steel plates. the interaction between adjacent core panels is characterized by a hard contact in the normal direction and by a frictionless contact in the tangential direction. concerning constraints, displacement orthogonal to the symmetry plane are prevented, the lower steel plate is constrained in the vertical direction on the bottom face and a vertical displacement δ is imposed on the loading line. the mesh is shown in the same figure: eight-node linear hexahedral elements (c3d8r continuum, 3-d, 8-node reduced integration) are used. 4.2. constitutive laws in this section, a description of the constitutive laws adopted for each material is proposed. the elastic phase of trc is defined introducing a poisson’s ratio of 0.2 and a young’s modulus of 30 gpa (according to literature results [17] on mortar characterized by equal maximum aggregate size and compressive strength). the plastic behaviour of trc is modelled through abaqus concrete damage plasticity model. no damage curve is inserted; hence, the model behaves as a plasticity model. in compression, an elasticperfectly plastic behaviour is assumed, imposing a yield strength of 97 mpa according to the values shown in section 2.3.1. the stress-irreversible strain relationship (table 3) adopted in tension is extracted from the experimental results of piquer specimens shown in figure 1a, considering strains computed basing on the measurements of displacement transducers applied on trc. trc tensile behaviour is assumed homogeneous over the layer thickness; this assumption is reliable according to [10]. point yield stress cracking strain t1 2.5 0 t2 5 0.00275 t3 20.81 0.017 table 3. trc plastic tensile constitutive law. concerning the core material, the elastic phase is defined introducing a young’s modulus of 20 mpa 101 i. g. colombo, m. colombo, m. di prisco acta polytechnica ctu proceedings figure 4. fe model: geometry and constraints (a) and validation of core material model considering three-point bending (b). figure 5. fe model results: numerical response compared with experimental results (a) and with easee bending behaviour (b); failure modes for specimen piquer2 showing core strut yielding at point c1 (c) and trc multicracking at point t3_sup (d). 102 vol. 33/2022 trc sandwich solution for energy retrofitting (according to experimental results in compression shown in figure 1b) and a poisson’s ratio of 0.2. the plastic behaviour of the core material is introduced using abaqus concrete damage plasticity model. also in this case, no damage curves are introduced. an elastic-perfectly plastic behaviour is assumed both in tension and in compression: a compressive yield strength of 0.30 mpa is used according to experimental results shown in figure 1b (average value for specimens of the same batch of core sample piquer2); a tensile yield strength of 0.053 mpa is adopted according to average experimental results of tensile tests performed on core specimens. with the aim of checking the validity of the material model adopted for the core, a three-point bending test performed on core prismatic specimen with size of 160×40×40mm3 was modelled. experimental and numerical curves are compared in figure 4b, demonstrating that the adopted material model is reliable. steel is assumed elastic with a young’s modulus of 210 gpa and a poisson’s ratio of 0.1. 4.3. numerical results the numerical response is shown in figure 5a in terms of load (p) vs. stroke. a good correlation is found with respect to the behaviour of specimen piquer2 that showed a hardening behaviour. a perfect overlapping is obtained in the initial phase of the test; the divergency of the two curves while the test is progressing could be related to a bond weakness at trc-core interfaces, leading to the final failure due to debonding already underlined in figure 3c. in order to investigate the failure mechanisms occurred for a specimen characterised by a good trccore bond, critical points of material constitutive laws are underlined on the numerical curve. in particular: t points refer to the trc layers (table 3), with subscript "inf" related to lower trc layer and subscript "sup" related to the upper trc layer; point c1 indicates the yielding of the strut in the core material. it is worth noting that the change in the slope of the global response corresponds to point c1 (figure 5c), when both trc layers are already cracked (figure 5d). in figure 5b the numerical response of piquer specimens, in which the core thickness was enhanced to 100 mm, is compared with the experimental behaviour of easee sandwich beams characterised by the same cross-section and size. 5. conclusions the bending behaviour of sandwich solutions developed within the easee and smart p.i.qu.e.r. projects was compared at lab-scale level. even though the response of the piquer sandwich beam showed lower maximum load and final displacement, a ductile hardening behaviour and a reasonable sustained load are guaranteed by the solution. it is worth noting that a good bond at trc-core interfaces must be guaranteed to obtain a proper behaviour of the composite. acknowledgements the research was developed within the smart p.i.qu.e.r. project, funded by regione lombardia (smart living; decree n. 14782 24/11/2017; id siage: 379284). authors would like to thank the partners; in particular, ipcbcnr and isoltech involved in the development of the core material. references [1] eur-lex. directive 2018/844/eu of the european parliament and of the council of 30 may 2018 amending directive 2010/31/eu on the energy performance of buildings and directive 2012/27/eu on energy efficiency, 2018. https://eur-lex.europa.eu/l egal-content/en/txt/?uri=celex:32018l0844. [2] european commission. data from eu building database, 2019. https: //ec.europa.eu/energy/en/eu-buildings-database. [3] j. hegger, m. horstmann. light-weight trc sandwich building envelopes. excellence in concrete construction through innovation 187-194, 2008. https://doi.org/10.1201/9780203883440.ch27. [4] a. einea, d. c. salmon, g. j. fogarasi, et al. state-of-the-art of precast concrete sandwich panels. pci journal 36(6):78-98, 1991. https://doi.org/10.15554/pcij.11011991.78.98. [5] a. shams, m. horstmann, j. hegger. experimental investigations on textile-reinforced concrete (trc) sandwich sections. composite structures 118:643-53, 2014. https: //doi.org/10.1016/j.compstruct.2014.07.056. [6] european commission. european project easee (2012-2016) envelope approach to improve sustainability and energy efficiency in existing multi-storey multi-owner residential buildings, grant no.: 285540, 2019. [7] i. g. colombo, m. colombo, m. di prisco. tensile behavior of textile reinforced concrete subjected to freezing-thawing cycles in un-cracked and cracked regimes. cement and concrete research 73:169-83, 2015. https: //doi.org/10.1016/j.cemconres.2015.03.001. [8] i. g. colombo, m. colombo, m. di prisco. bending behaviour of textile reinforced concrete sandwich beams. construction and building materials 95:675-85, 2015. https: //doi.org/10.1016/j.conbuildmat.2015.07.169. [9] i. g. colombo, m. colombo, m. di prisco. trc precast façade sandwich panel for energy retrofitting of existing buildings. aci sp. pub. sp 305, 2015. https://doi.org/10.14359/51688590. [10] i. g. colombo, m. colombo, m. di prisco, et al. analytical and numerical prediction of the bending behaviour of textile reinforced concrete sandwich beams. journal of building engineering 17:183-95, 2018. https://doi.org/10.1016/j.jobe.2018.02.012. 103 https://eur-lex.europa.eu/legal-content/en/txt/?uri=celex:32018l0844 https://ec.europa.eu/energy/en/eu-buildings-database https://doi.org/10.1201/9780203883440.ch27 https://doi.org/10.15554/pcij.11011991.78.98 https://doi.org/10.1016/j.compstruct.2014.07.056 https://doi.org/10.1016/j.cemconres.2015.03.001 https://doi.org/10.1016/j.conbuildmat.2015.07.169 https://doi.org/10.14359/51688590 https://doi.org/10.1016/j.jobe.2018.02.012 i. g. colombo, m. colombo, m. di prisco acta polytechnica ctu proceedings [11] i. g. colombo, m. colombo, m. di prisco, et al. trc sandwich panel for energy retrofitting exposed to environmental loading. aci sp. pub. sp 326, 2018. https://doi.org/10.14359/51711059. [12] i. g. colombo, m. colombo, m. di prisco. precast trc sandwich panels for energy retrofitting of existing residential buildings: full-scale testing and modelling. materials and structures 52(5), 2019. https://doi.org/10.1617/s11527-019-1406-1. [13] g. salvalai, m. m. sesana, g. iannaccone. deep renovation of multi-storey multi-owner existing residential buildings: a pilot case study in italy. energy and buildings 148:23-36, 2017. https://doi.org/10.1016/j.enbuild.2017.05.011. [14] i. g. colombo, m. colombo, m. di prisco, et al. lightweight trc sandwich panels with sustainable diatomite-based core for energy retrofitting of existing buildings. advances in building energy research 15(2):231-52, 2019. https://doi.org/10.1080/17512549.2019.1697752. [15] en 196-1:2005. standard for mortar. methods of testing cement part 1: determination of strength, 2015. [16] l. verdolotti, b. liguori, i. capasso, et al. synergistic effect of vegetable protein and silicon addition on geopolymeric foams properties. journal of materials science 50(6):2459-66, 2014. https://doi.org/10.1007/s10853-014-8801-3. [17] w. brameshuber. report of rilem technical committee 201-trc rilem publications, 2016. 104 https://doi.org/10.14359/51711059 https://doi.org/10.1617/s11527-019-1406-1 https://doi.org/10.1016/j.enbuild.2017.05.011 https://doi.org/10.1080/17512549.2019.1697752 https://doi.org/10.1007/s10853-014-8801-3 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0222 acta polytechnica ctu proceedings 38:222–227, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague natural materials in building construction – annual evaluation daniela michálková∗, pavol ďurica university of žilina, faculty of civil engineering, department of building engineering and urban planning, univerzitná 8215/1, 010 26 žilina, slovakia ∗ corresponding author: daniela.michalkova@uniza.sk abstract. the construction industry’s focus on a low-carbon economy will result in the need for a deeper examination of natural-based building materials. from an environmental point of view, the benefits of these materials are undeniable. however, it is necessary to consider their shortcomings in other areas of design in terms of building thermal engineering. this article observes and evaluates a wall designed for a wooden building with almost zero energy demand in year-round operation and subsequent assessment in confrontation with a different composition, seemingly more advantageous in thermal resistance and humidity regime. these assemblies are under long-term examination within the pavilion research of the authors’ workplace, in laboratory conditions from the interior side. at the same time, they are exposed to the realistic boundary conditions of the external environment. the paper includes an environmental assessment of two compositions, variating in the used material. the research shows that the wall composition of natural materials is more advantageous from an ecological perspective and can also show favourable effects in terms of temperature and relative humidity regulation. keywords: timber-framed, natural building materials, temperature, relative humidity, environmental assessment. 1. introduction the acute need to improve the environment is an increasingly trending topic. sustainable development includes many issues, among others, improvement of the building industry [1]. in order to maintain liveable conditions of the earth, it is necessary to reduce the use of materials with great primary energy demand [2]. according to the paris agreement [3], the united nations – 191 countries, which signed the document – are bound to keep the global temperature rise this century below 2 °c and report every five years their actions to ensure the goal. the harmonised requirements for construction products summarises the eu regulation no. 305/2011 of the european parliament and council [4], which significantly supports the use of raw and secondary materials in the building industry. at the same time, the requirements of current legislation regarding the thermal protection of buildings [5] are equally important. 2. experimental wall samples this article aims to present two different wall assemblies containing various materials. both face south orientation with 15 ° inclinations to the west, and both are exposed to the natural exterior climate, measured on the laboratory roof. the interior space uses air conditioning, and its’ temperature and relative humidity are appropriately measured. the research contains overall ten wall samples. all of them are under long-term investigation to monitor the temperature and relative humidity within the timber-framed multi-layered constructions. the monitoring is in three high levels – 0.5 m under the ceiling, in the middle of the structure height, and 0.5 m above the floor – in all material interfaces. in this paper, we present values in the middle of structures’ height, apart from the interface between phenolic foam and osb in s1, bearing in mind that we would have nothing to compare these values to in wall s1. used are ntc thermistors with the accuracy of ± 0.2 °c for temperature and capacity probes with the precision of ± 2 % for the relative humidity. the first wall in this paper, marked as s1 (south 1), is the only one with solely natural materials. the second, marked as s2 (south 2), was chosen for comparison, as it is the most similar to the first assembly. both these wall constructions are diffusely open, enabling air and water penetration and thereby reducing the risk of fungi and mould settling. wall s1 in figure 1 solely consists of timber frame, timber log profiles from exterior and interior and sheep wool within. s2 has the same exterior layer, followed by basalt fibre thermal insulation isover granulate, phenolic foam insulation kingspan kooltherm k5, finished by osb. the probes are in three depths within the wall: inside under the log profile, in the middle of thermal insulation, between wool and interior log profile (s1), and between isover granulate and kingspan kooltherm (s2). table 1 comprises the materials of both wall assemblies with tier main physical characteristics and environmental indicators. among them is d – thickness, 222 https://doi.org/10.14311/app.2022.38.0222 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 natural materials in building construction – annual evaluation figure 1. left south wall 1 (s1), right south wall 2 (s2) – layers, probes placement. material d λ µ ρ c gw p p ei ap [m] [ w m3·k ] [-] [kg/m3] [ j kg·k ] [co2eq/kg] [mj/kg] [so2eq/kg] timber profile [6] 0.068 0.180 157 400 2510 0.109 1.959 0.00128 sheep wool [7] 0.220 0.042 1.5 16 1720 0.537 19.324 0.00463 basalt fibre ti [6] 0.220 0.040 1 50 1020 0.346 21.363 0.01413 phenolic foam ti [8] 0.040 0.021 35 35 1400 3.821 96.515 0.01742 osb [6] 0.012 0.130 50 650 1700 0.481 12.506 0.00210 table 1. layer materials of both assemblies and their main characteristics. figure 2. s1 environmental evaluation (baubook.at). λ – thermal conductivity factor, µ – water vapour diffusion factor, ρ – bulk density, c – specific heat capacity, gw p – global warming potential, p ei – primary energy intensity, and ap – acidification potential, listed. ti stands for thermal insulation. figure 3. s2 environmental evaluation (baubook.at). 3. environmental assessment the environmental benefits of the first wall are undeniable. on the other hand, the second wall consists of more durable materials and therefore may balance the negatives. thus rose the need to establish the environmental impact of both structures. the assessment stems from the baubook website [9], which is 223 daniela michálková, pavol ďurica acta polytechnica ctu proceedings figure 4. comparison of the temperature in the winter period. figure 5. comparison of the relative humidity in the winter period. an online platform that promotes sustainable building design. this website enables the evaluation of the building structures in terms of the oi – ecology index (ökoindex) [10] that considers the environmental indicators such as gw p , ap , and p ei. for comparison, the standard value for common wall assemblies is 70 pts/m2 [11]. the results of the environmental assessment are in figure 2 and figure 3. for comparison, the u -value (heat transfer coefficient) is according to stn en iso 6946 [12] 0.196 w/(m2 · k) for s1 and 0.145 w/(m2 · k) for s2, calculated as inhomogeneous wall structure. as shown, the wall assembly marked as s1 consisting of timber log profiles and sheep wool is classified as a++ according to oi-class, with ∆oi3 value 6 pts/m2. the structure s2 of timber log profile, basalt fibre insulation, phenolic foam insulation and oriented strand board reached value 38 pts/m2 and thus fell to the a+ class. 4. temperature and relative humidity the temperature and relative humidity, measured in three depths of the wall (figure 1), provided large datasets throughout the year. to simplify the comparison, these quantities are in this paper divided into three sections – winter period from 20th january to 2nd february, spring period from 19th march to 1st april, and summer period from 18th june to 1st july. dotted lines present the temperature and relative humidity of exterior air. interior climate was set to 20 °c temperature and 50 % relative humidity. the values in placement s1.1 or s2.1 are further referred to as the interior. in this case, we use this term only to simplify the text slightly. however, we would like to emphasise the layers between the probes placement and the interior, which we are aware of. 4.1. winter period the graph in figure 4 shows the temperature in both structures during two weeks of winter. the yellow dotted line represents the exterior air temperature. the other six lines stand for the temperature, whereas the different placements have different colours. the wall assemblies are distinguished by colour shade. although the temperature of s1 is from the exterior and in the middle of the wall lower than in s2, it reaches almost 5 °c higher values from the interior. that is naturally caused by the phenolic foam from the interior in the case of s2. however, s2 is ex224 vol. 38/2022 natural materials in building construction – annual evaluation figure 6. comparison of the temperature in the spring period. figure 7. comparison of the relative humidity in the spring period. pected to reach a lower temperature under the exterior cladding due to the higher thermal performance. another investigation consequently followed this analysis. unfortunately, we found out that in s2 was created an air cavity, causing an additional thermal bridge. significant is the alignment during the days where the amplitude of exterior temperature was smaller. the temperature these days was in the middle almost identical. figure 5 displays relative humidity in the same period. the walls and probe placement is distinguished the same way as in the case of temperature. the regulation of relative humidity is, in the case of s1, very significant. the difference between exterior and interior relative humidity is almost 50 %, whereas in s2, only 20 %. it may sound unreasonable, but the explanation behind this statement is the following. both assemblies are under constant interior boundary conditions, which means we can see only the influence of exterior climate. although s1 reaches higher relative humidity from the outside, it meets the values of s2 relative humidity already in the middle of the insulation layer. the sheep wool was able to further lower the relative humidity to circa 30 %, whereas in basalt fibre, the value is around 40–50 %. this is the reason why we claim that sheep wool has the ability to regulate relative humidity better. the relative humidity in the middle is very similar in both assemblies. the variations occur under the outside and inside layers of both. also interesting is the course of the humidity itself. the wall s1 reached steadier values throughout the whole period. 4.2. spring period to represent the spring period, we have selected two weeks from 19th march to 1st april. figure 6 shows the temperature comparison. the daily amplitude is naturally more significant. from the outside, the solar impact on the temperature of the exterior layer is evident, where the temperature in the walls is 15 °c higher than the surrounding air temperature. other than that, the course of temperature is similar to the winter period, where the middle temperature is almost identical, and the interior temperature differs the most. figure 7 presents the relative humidity. the difference between exterior and interior humidity is here not so significant and gets even smaller towards the end of selected weeks. however, the internal relative humidity is again lower in the case of s1 compared to the wall s2, creating a difference of nearly 10 %. 225 daniela michálková, pavol ďurica acta polytechnica ctu proceedings figure 8. comparison of the temperature in the summer period. figure 9. comparison of the relative humidity in the summer period. 4.3. summer period the summer period is represented by 14 days from 18th june to 1st july. as expected, the temperature amplitude is large, with a maximum variance of 24 °c. it is displayed in figure 8, where a noticeable temperature shift is apparent, in some cases reaching 6 hours. according to figure 8, wall s1 is capable of higher temperature regulation. to support this statement, notice that the temperature from the outside is, in its case, higher than in s2. in contrast, already in the middle of thermal insulation is vice versa, where the s1 temperature is 3 °c lower than the s2. figure 9 shows the course of relative humidity during the summer period. in this case, the difference is foremost in the interior layer, caused by the air relative humidity rise due to some technical difficulties with the air conditioning. however, the relative humidity in the middle of both walls is relatively consistent, varying from 50 to 60 %. 5. conclusion although both wall assemblies appear very similar, their difference is rather significant. the first structure consists solely of natural materials with minor adjustments to enable their use in buildings. the second wall uses a combination of natural materials – timber frame, exterior timber log profile, and oriented strand board – and synthetic materials, such as basalt fibre thermal insulation and phenolic foam to improve its thermal resistance. nevertheless, the difference is not significant enough to balance the environmental impact. as stated in the section 3 (environmental assessment), both structures fall into the highest categories in terms of ecoindex. however, wall s2 requires almost double primary energy in comparison with s1. moreover, the acidification and global warming potential are also double. in terms of ∆oi3 is for wall s2 is more than six times greater – 38 pts/m2 versus 6 pts/m2 for s1. from the perspective of used materials, wall s2 is expected to show better temperature and relative humidity results. the opposite turned out to be the case, where the first assembly showed its capability of not only competing with s2 in terms of thermal performance but, in some cases, even outstanding it. wall s1 provides more significant temperature decrement towards the interior, obvious foremost in the summer period. the regulation of relative humidity within s1 was in some cases 30 % greater opposite to 226 vol. 38/2022 natural materials in building construction – annual evaluation wall s2. however, it is needless to say that neither of the wall assemblies showed relative humidity values that could initialise the growth of moulds and fungi or cause material degradation. the main conclusion of this paper is not to underestimate such wall assemblies, which can prove their worth not only in terms of environmental sustainability but also in terms of overall performance. future research could contribute by numerical simulation of both wall structures within an actual building and deeper evaluation from an lca point of view. acknowledgements the research of this paper was enabled through the vega project nr. 1/0673/20 theoretical and experimental analysis of energy effective and environmentally friendly building envelopes. references [1] d. a. suhamad, s. p. martana. sustainable building materials. iop conference series: materials science and engineering 879(1):012146, 2020. https://doi.org/10.1088/1757-899x/879/1/012146 [2] w. s. e. ismaeel. material selection for sustainable buildings. in conference paper arcom 2021. 2021. [3] united nations. paris agreement to the united nations framework convention on climate change, 2015. [2022-02-28]. https://unfccc.int/sites/ default/files/english_paris_agreement.pdf [4] european union. regulation (eu) no 305/2011 of the european parliament and of the council of 9 march 2011 laying down harmonised conditions for the marketing of construction products and repealing council directive 89/106/eec. [2022-02-28]. http://data.europa.eu/eli/reg/2011/305/oj [5] stn 73 0540-2+z1+z2 thermal protection of buildings. thermal performance of buildings and components. part 2: functional requirements, 2019. [6] inštitút pre energeticky pasívne domy. ekologické materiály pre energeticky pasívne domy (epd), 2020. [2022-02-28]. https://ozartur.sk/wpcontent/uploads/2020/08/iepd-materialy.pdf [7] u-values for sheep wool insulation, 2014. [2022-02-28]. http://www.isolena.cz/wp-content/uploads/0259_ iso17_architektenordner_2_u-werteproduktdatenbl%c3%a4tterakustikd%c3%a4mmung_cz_iii.pdf [8] kingspan. environmental product declaration – kingspan kooltherm k5, 2019. [2022-02-28]. https://www.kingspan.com/pl/plpl/produkty/izolacje/pliki-dopobrania/certyfikaty/epd-kooltherm-k5 [9] baubook component calculator, 2018. [2022-02-28]. https://www.baubook.at/btr/ [10] oekoindex component evaluation, guideline for the ecological evaluation of component constructions by means of eco-index v1.4„ energy institute vorarlberg, dornbirn. [11] a. iringová. design of envelopes for timber buildings in terms of sustainable development in the low-energy construction. iop conference series: materials science and engineering 415:012010, 2018. https://doi.org/10.1088/1757-899x/415/1/012010 [12] stn en iso 6946 building components and building elements. thermal resistance and thermal transmittance. calculation methods, 2001. 227 https://doi.org/10.1088/1757-899x/879/1/012146 https://unfccc.int/sites/default/files/english_paris_agreement.pdf https://unfccc.int/sites/default/files/english_paris_agreement.pdf http://data.europa.eu/eli/reg/2011/305/oj https://ozartur.sk/wp-content/uploads/2020/08/iepd-materialy.pdf https://ozartur.sk/wp-content/uploads/2020/08/iepd-materialy.pdf http://www.isolena.cz/wp-content/uploads/0259_iso17_architektenordner_2_u-werte-produktdatenbl%c3%a4tter-akustikd%c3%a4mmung_cz_iii.pdf http://www.isolena.cz/wp-content/uploads/0259_iso17_architektenordner_2_u-werte-produktdatenbl%c3%a4tter-akustikd%c3%a4mmung_cz_iii.pdf http://www.isolena.cz/wp-content/uploads/0259_iso17_architektenordner_2_u-werte-produktdatenbl%c3%a4tter-akustikd%c3%a4mmung_cz_iii.pdf http://www.isolena.cz/wp-content/uploads/0259_iso17_architektenordner_2_u-werte-produktdatenbl%c3%a4tter-akustikd%c3%a4mmung_cz_iii.pdf https://www.kingspan.com/pl/pl-pl/produkty/izolacje/pliki-do-pobrania/certyfikaty/epd-kooltherm-k5 https://www.kingspan.com/pl/pl-pl/produkty/izolacje/pliki-do-pobrania/certyfikaty/epd-kooltherm-k5 https://www.kingspan.com/pl/pl-pl/produkty/izolacje/pliki-do-pobrania/certyfikaty/epd-kooltherm-k5 https://www.baubook.at/btr/ https://doi.org/10.1088/1757-899x/415/1/012010 acta polytechnica ctu proceedings 38:222–227, 2022 1 introduction 2 experimental wall samples 3 environmental assessment 4 temperature and relative humidity 4.1 winter period 4.2 spring period 4.3 summer period 5 conclusion acknowledgements references my title https://doi.org/10.14311/app.2022.36.0090 acta polytechnica ctu proceedings 36:90–98, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague influence of the structural integrity management on the levelized cost of energy of offshore wind: a parametric sensitivity analysis marko kinnea, ∗, muhammad farhana, ronald schneidera, sebastian thönsa, b a bundesanstalt für materialforschung und -prüfung (bam), unter den eichen 87, 12205 berlin, germany b lund university, faculty of engineering, division of structural engineering, department of building & environmental technology, faculty of engineering, box 118, 221 00 lund, sweden ∗ corresponding author: marko.kinne@bam.de abstract. the levelized cost of energy (lcoe) is an important measure to quantify the macro-economic efficiency of an offshore wind farm and to enable a quantitative comparison with other types of energy production. the costs of the structural integrity management which is required to ensure an adequate lifetime reliability of the turbine support structures are part of the operational expenditures of an offshore wind farm. an optimization of the structural integrity management may reduce the operational expenditures and consequently the lcoe. however, the effect of the structural integrity management on the lcoe is hardly known. to investigate this effect, this paper presents a sensitivity analysis of the lcoe of a generic offshore wind farm. the probabilistic models of the parameters influencing the lcoe are based on a literature study including an explicit model for the structural integrity management. the analysis reveals that lcoe may potentially be reduced if an optimization of the structural integrity management enables a service life extension. keywords: levelized cost of energy, sensitivity analysis, structural integrity management. 1. introduction in the context of the energy transition in germany and the european union, the relevance of renewable energy is steadily increasing in contrast to conventional energy sources such as coal-fired power plants. besides hydropower and solar energy, offshore and onshore wind energy is becoming a major part of renewables energies and is likely to be expanded in the future. over the past 20 years, the total number of wind turbines increases from 9400 to 29600 in germany, whereof 1500 wind turbines with a total capacity of 7.7 gw are currently installed offshore [1, 2]. according to the eu strategy presented by the eu commission in 2020, €800 billion will be invested in offshore renewable energy over the next 30 years to increase the capacity of offshore wind energy from 12 gw to 60 gw by 2030 with an aim to reach a total capacity of 300 gw by 2050 [3]. in addition to the investments in the expansion of renewable energy, however, the efficiency of renewable energy sources measured in terms of the levelized cost of energy (lcoe) must also be considered to ensure that these energy sources are competitive. as an example, in 2018 the lcoe of offshore wind energy in germany was still higher than the lcoe of other main energy carrier such as brown coal [4]. besides the reduction of the investment costs and the increase of the power production, wind farm operators optimize structural integrity management (sim) procedures to reduce the lcoe. such optimizations are constrained by different rules and standards (e.g. requirements set out in germany by the bundesamt für seeschifffahrt und hydrographie (bsh) or in united kingdom by the department for business, energy & industrial strategy and the marine management organisation). the challenge is to optimize the operation of an offshore wind farm including the sim in compliance with the governing rules and standards. in the existing literature, the overall costs related to the sim are low in contrast to other costs such as the investment costs and the costs related to the turbine integrity management [5, 6]. based on this information, the question is whether it is worthwhile to optimize the sim for the support structures in an offshore wind farm to reduce the lcoe. in this paper, a sensitivity analysis is performed to quantify the influence of the sim on the lcoe of a generic offshore wind farm and to investigate the opportunity of upgrading/optimizing the sim to reduce the lcoe (i.e., an optimization of the inspection and maintenance strategy as well as monitoring systems). to investigate the influence of the sim on the lcoe, the operational expenditures are divided in a part related to the structures and a part related to the turbines. the lcoe is calculated on the basis of current scientific literature, in which the operational expenditures related to the support structures in an offshore 90 https://doi.org/10.14311/app.2022.36.0090 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 36/2022 levelized cost of energy of offshore wind figure 1. illustration of a global sensitivity analysis, adapted from [5]. wind farm are determined based on the requirements defined by the bsh. to determine the effect of the sim and other influencing parameters, the first order sensitivity index (sobol index) is calculated based on a parametric model of the lcoe. 2. levelized cost of energy the lcoe quantifies the average net costs for generating electricity with a certain type power plant. it is defined as the ratio between (a) the sum of the costs consisting of the capital expenditures (capex) and accumulated operational expenditures (opex) and (b) the sum of the annual energy production (aep) over the service life l (equation 1) [6]. lcoe = cap ex + l∑ t=1 op ex (1 + i)t l∑ t=1 aep (1 + i)t (1) the capex contain the investment costs of an offshore wind farm, while the opex include the costs related to the operating of the support structures and turbines including the costs for monitoring, inspection and maintenance. opex and aep are discounted based on the discount rate i. the aep of an offshore wind farm is computed as the product of the nominal (turbine) capacity nomcap, the nominal capacity availability factor nomcapava, the turbine availability turbavaf ac, the number of wind turbines nwt and the feed in tariff f eedin ((equation 2) [7]. aep = nwt · nomcap · 365.25 · 24 · nomcapava · tubravaf ac · f eedin (2) the capital expenditures are calculated in function of the number of wind turbines nwt in a particular wind farm and the investment costs per wind turbine turbinvest (equation 4) cap ex = nwt · turbinvest (3) the operational expenditures depend on the number of wind turbines nwt and the operational costs per wind turbine turboperation (equation ??). op ex = nwt · turboperation (4) 3. sensitivity analysis a sensitivity analysis provides information on how input parameters x and their uncertainties influence the output y of a deterministic model. a distinction is made between local and global sensitivity analyses [5]. a local sensitivity analysis studies the influence of variability of the input parameters on the model output around a point x0. generally, the influence of small changes in input parameters on the model output is investigated. a global sensitivity analysis determines the influence of the input parameters by varying them over their entire domain. the basic procedure of a global sensitivity analysis is illustrated in figure 1. the input parameters may have the same or different marginal probabilistic distributions. the variance of the model output y depends on the variance of the input parameters and the relations implemented in the deterministic model. in a global sensitivity analysis, the share vi of the variance of model output y that is caused by the input parameter xi is determined. a sensitivity analysis can pursue different goals [8, 9]: • robustness: understanding the model’s robustness with regard to the input parameters • ranking: rank input factors according to their importance • screening: identifying input factors with minor importance, which can be omitted • mapping: identifying the areas of input parameters, which lead to extreme model outputs • decision support: quantify the effect of decision variables on the model output. in this contribution, the sensitivity analysis is performed to rank the parameters influencing the lcoe according to their importance. to this end, the variance-based sensitivity analysis is applied which defines an input parameter’s importance in terms of its contribution to the variance of the model output v ar[y ] = v ar[g(x)]. in this approach, the influence of the uncertainty in an input parameter xi on the variance v ar[y ] is quantified by a first order measure vi (equation 5) [9] vi = v arxi {ex−i [ g (x) |xi]} (5) 91 m. kinne, m. farhan, r. schneider, s. thöns acta polytechnica ctu proceedings wind turbine cap exstruc/k€/mw cap exturbine/k€/mw reference 4-a-14 677 2021 [10] 4-d-14 861 2131 [10] 8-a-14 689 1997 [10] 8-d-14 722 2180 [10] 6.1-mw fixed-bottom offshore wind turbine 723 2823 [11] generic offshore wind turbine 1065 2891 [12] table 1. capex of different offshore wind turbines divided into a structural and a turbine part. in which ex−i [ g (x) |xi] is the expected value of y = g(x) with respect to all input variables except xi, which is fixed. vi is zero, if xi has no effect on y and the expected values is constant with regards to xi. if xi is the only random variable with an effect to the model output, vi is equal to the full variance of the model output v ar[y ]. the first order measure vi can be estimated using a monte carlo simulation, which results in nm cs samples of the model output [9]. for every input parameter xi, the sample pairs (xi, y ) are ordered according to the value of xi. the samples are divided into blocks of size nb and the mean value µyi of each block is an estimate of ex−i [ g (x) |xi]. the variance of the block means µyi is an estimate of vi. a more efficient way for determining the first order measure vi is the sobol sequence [13]. based on vi, the first order sensitivity index si can be computed as follows [14] (equation 6). si = vi v ar[g(x)] = v arxi {ex−i [ g (x) |xi]} v ar[g(x)] (6) 4. numerical study to study the influence of the structural integrity management on the lcoe, the capital and operational expenditures are divided into a part related to the structure and a part related the turbine (equation 7 and 8). cap ex = nwt · (cap exstruc + cap exturbine) (7) op ex = nwt · (op exstruc + op exturbine) (8) the capital expenditures cap exstruc and cap exturbine for one offshore wind turbine are obtained from the data listed in table 1. the values provided in table 1 are normalized by the nominal turbine capacity. based on the data listed in table 1, a range of [677, 1065]/k€/mw is assumed for the capital expenditures related to the structure cap exstruc, while the capital expenditures related to the turbine cap exturbine exhibit a range of [1997, 2891]/k€/mw. the operational expenditures related to the structure are derived based on a study documented in [15]. the characteristics as well as the environmental and operational conditions of the wind farm considered in [15] are summarized in table 2. based on this wind farm, the operational expenditures are estimated for an inspection and monitoring strategy with three different settings (optimistic, average, pessimistic) [15]. the inspection and monitoring strategy is in accordance with the requirements defined by the bsh as summarized in table 3. according to the requirements set out by the bsh, general visual inspection (gvi) of the primary and secondary steelwork of the support structures in the wind farm above water is performed every year to provide a general overview on any obvious mechanical damage, fatigue or corrosion. typically, such an inspection is performed by inspectors from a crew transfer vessel. in addition, 25% of the support structure are inspected using close visual inspection (cvi) and detailed visual inspection (dvi). the aim of cvi is to identify fatigue or corrosion and determine whether non-destructive testing (ndt) would be necessary to investigate the welded connections. in contrast to gvi, such an inspection is carried out closer to the support structure. according to [15], the objective of dvi is to determine the extent of detected damage. for this purpose. methods of non-destructive testing (ndt) may also be used. below water, the number of support structures considered for gvi is reduced to 25% every year, while the aim of obtaining a general overall overview on the condition of the support structures remains the same. however, the difference is that a remotely operated vehicle is used to perform the inspection. cvi below water is performed to identify corrosion or fatigue damage and to determine whether ndt would be necessary to inspect the structural components [15]. for this type of inspection, good environmental conditions and visibility as well as marine growth cleaning are required. dvi is used to determine the extent of detected damage using ndt methods [15]. the number of support structures inspected by cvi and dvi per 92 vol. 36/2022 levelized cost of energy of offshore wind characteristic unit value number of owts − 100 turbine capacity mw 5 wf area km2 50 average distance to port km 50 average water depth m 30 foundation type − monopile number of offshore substations − 1 average wind speed m/s 10 (at hub height) tidal conditions s 0.5 (hat to lat) 50-year wave m 6.5 current m/s 1 number of export cables − 1 table 2. wind farm characteristics and environmental and operational conditions [15]. activity shms in 10% of wts inspection frequencyduring service life above water gvi of primary and secondary steelwork 100% every year 25 cvi of primary and secondary steelwork 25% every year 6.25 dvi of primary and secondary steelwork 25% every year 6.25 seabed scour survey 100% the 2 first yearsand then 25% every year 7.75 subsea marine growth survey 25% every year 6.25 cathodic protection potential survey 100% the 2 first years and then 25% every year 7.75 cvi of the grouted connection 25% every year 6.25 below water gvi of primary and secondary steelwork 25% every year 6.25 cvi of primary and secondary steelwork 25% every year 6.25 dvi of primary and secondary steelwork 25% every year 6.25 table 3. inspection strategy [gvi: general visual inspection; cvi: close visual inspection; dvi: detailed visual inspection] [15]. boundary conditions optimistic average pessimistic op exstruc 1.2% 1.6% 1.9% op exturbine 98.8% 98.4% 98.1% table 4. relative contribution of op exturbine and op exstruc to the op ex. 93 m. kinne, m. farhan, r. schneider, s. thöns acta polytechnica ctu proceedings figure 2. illustration of the distribution types for a given value range defined in terms of the min. and max. value. year below water is the same as the number of support structures inspected by cvi and dvi above water. a seabed scour survey, a subsea marine growth survey and an inspection of the cathodic protection are performed for 100% of the support structures in the first two years. thereafter, this inspect effort is reduced to 25% of the support structures in each year. the objective of the marine growth survey is to determine the coverage, thickness and type of marine growth on the support structures and sacrificial anodes to guide decisions on removing the marine growth. in a scour survey, the seabed around the support structures is inspected to monitor changes in the topology (local and global). an inspection of the cathodic protection determines if there is adequate global cathodic protection of the submerged part of the support structures. according to the requirements defined by the bsh, continuous sensor-based monitoring systems (structural health monitoring systems and conditional monitoring systems) have to be installed on 10% of the wind turbines. the operational expenditures related to the structures op exstruc and turbines op exturbine relative to the opex are given in table 1. these values are determined in [15] based on the inspection and monitoring scenario described above in conjunction with optimistic, average and pessimistic estimates of the associated costs. from table 4 it can be seen that the op exstruc ranges between 1.2% to 1.9% of the overall opex. the accumulated discounted operational expenditures l∑ t=1 op exstruc + op exturbine (1 + i)t of an offshore wind farm correspond to approximately 25% of the total costs consisting of capital and operational expenditures [21]. to estimate the ranges of the absolute values (min./max. values) of op exturbine and op exstruc, a simple monte carlo simulation is performed. in this analysis, the capex is modelled as a uniform distributed random variable. furthermore, the service life and the discount rate are also modelled as uniform distributed random variables. in combination with the knowledge on the relative values of op exturbine and op exstruc (modelled as uniform distributed random variables) (table 4), the ranges of the absolute values of op exturbine and op exstruc can be estimated. from this analysis it was found that the operational expenditures op exturbine exhibit a range of [64, 135]/k€/mw/year, while the operational expenditures op exstruc are [0.8, 2.6]/k€/mw/year. in the current case study, it is assumed that the service life l of the wind farms may be extended from 20 years up to 30 years [20]. the nominal capacity availability factor nomcap ava describes the ratio between the average power output of a wind turbine (usually one year) to its maximum output and exhibits a range of [0.4, 0.5] [16]. this factor is affected by a variety of influencing factors and conditions including the windspeed. the turbine availability factor tubrava f ac specifies the expected average availability over the service life. considering only downtime the turbine itself. the factor accounts for the loss of energy associated with amount of time, when the turbine is not available to 94 vol. 36/2022 levelized cost of energy of offshore wind variable unit uniformdistribution normal distribution ref. lower bound upper bound mean value standard deviation (99.7%) standard deviation (90%) number of wind turbines nwt 100 100 100 0 0 nominal capacity nomcap [mw] 6 6 6 0 0 nominal capacity availability factor nomcap ava 0.4 0.5 0.45 0.02 0.03 [16] turbine availability factor turbava f ac 0.857 0.996 0.93 0.025 0.04 [17] feed in tariff f eedin [€/kwh] 0.14 0.16 0.15 0.003 0.006 [18] turbine invest: structure cap exstruc [k€/mw] 667 1065 871 70 118 table 1 turbine invest: turbine cap exturbine [k€/mw] 1997 2891 2444 162 271 table 1 operational expenditures: structure cap exstruc [k€/mw/year] 0.8 2.6 1.7 0.3 0.54 [5, 17] operational expenditures: turbine part cap exturbine [k€/mw/year] 64 135 99.5 12.2 21.6 [5, 17] discount rate i 0.06 0.08 0.07 0.003 0.007 [19] service life t [year] 20 30 25 1.8 3 [20] table 5. parameters of the probabilistic models applied in the sensitivity study. produce energy, e.g. due to failure or maintenance. the availability of the turbine per year is in the range of [0.857, 0.996] [17]. the feed in tariff f eedin describes the state fixed remuneration for electricity to subsidize certain types of electricity production, e.g. electricity generated by offshore wind. it is in the range of [0.14, 0.16]/€/kwh for offshore wind [18]. the discount rate i is in the range of [0.06, 0.0825]/year for offshore wind in germany [19]. often, only the minimal and maximal value of the parameters are available in the literature, but not the distribution type. first, it is assumed that the parameters influencing the lcoe are normal distributed. the calculation of the sensitivity indexes is performed for two different standard deviations of the input parameters. it is assumed that the input parameter ranges account for 99.7% and 90% of the values, whereby both ranges are symmetric to the mean value. in addition, the sensitivity indexes are also calculated 95 m. kinne, m. farhan, r. schneider, s. thöns acta polytechnica ctu proceedings figure 3. first order sensitivity indices of the parameters influencing the lcoe. parameter first order sensitivity index normal distributed parameter range: 90% normal distributed parameter range: 99.7% normal distributed parameter range: 90% and nomcap ava as weibull distributed uniform distributed nominal capacity availability factor 0.25 0.25 0.54 0.25 turbine availability factor 0.11 0.12 0.08 0.12 feed in tariff 0.09 0.09 0.06 0.09 capital expenditures: structure 0.04 0.05 0.03 0.05 capital expenditures: turbine 0.2 0.2 0.13 0.2 operational expenditures: structure 0.01 0.01 0.01 0.01 operational expenditures: turbine 0.16 0.17 0.11 0.16 discount rate 0.11 0.11 0.07 0.11 service life 0.09 0.08 0.05 0.08 table 6. first order sensitivity indices of the different input parameters. for uniform distributed input parameters (figure 2). moreover, in an additional calculation, the nominal capacity availability factor is modelled as a weibull distributed random variable to reflect the large influence of the wind speed [22]. the parameters of the weibull distribution are determined by assuming that 5% of the values are smaller than 0.4 (the lower value of the parameter range) and 5% are larger than 0.5 (the upper value of the parameter range). this results in a scale parameter of 0.47 and a shape parameter of 18.23. max. value in the current case study, realizations 96 vol. 36/2022 levelized cost of energy of offshore wind of the turbine availability factor tubrava f ac which are larger than one are set to be equal to one. note that, in the current example, the probability that the nominal capacity availability factor nomcap ava is smaller than zero or larger than one is zero regardless of whether it is modelled as a weibull or normal distributed random variable. in the calculation of the sensitivity indices, several simplifications are made. first, the input parameters are modelled as independent random variables. however, in reality dependence among some of the parameters exist, e.g. between the turbine availability factor and the costs due to maintenance / repair or investment costs. another simplification is made in that the lcoe is determined without considering failure of the support structures, which are typically associated with large costs. this simplification is reflected in the model of the op exstruc, which describes only the inspection and monitoring costs without taking into account any failure costs. further simplifications are made in the probabilistic modelling of the input parameters. as an example, the feed in tariff f eedin and the discount rate i are modelled as time-invariant random variables even though they typically vary with time. the influence of the adopted simplifications has to be investigated in future research. the parameters of the probabilistic models applied the sensitive analysis of the lcoe are provided in table 5. according to figure 3 and table 6, the indices depend on the distribution types. when modelling all the input parameters with the same distribution type (normal or uniform distributed), almost the same sensitivity indices are obtained. when modelling the nominal capacity availability factor as a weibull distributed random variable, its influence increases, while the influence of the remaining parameters decreases. however, the trend in the relative importance of the parameters remains approximately the same. it can be seen that the nominal capacity availability factor has the highest (first-order) sensitivity index, followed by the capital and operational expenditures related to the turbines. the discount rate, the turbine availability factor, and the feed in tariff have sensitivity indices in the same order of magnitude. the operational expenditures related to the support structures have always the smallest sensitivity index. the sensitivity index of the capital expenditures related to the support structure is smaller than the sensitivity index of the service life. 5. summary and concluding remarks in this contribution, a sensitivity analysis is performed to quantify the influence of the structural integrity management (sim) on the levelized cost of energy (lcoe) of an offshore wind farm. the lcoe of a wind turbine/ farm describe the average net present cost, which arise from the conversion from wind to electrical power. to quantify the influence of the structural integrity management, the operational expenditures are divided into a part related to the structures and a part related to the turbines. the input parameters and their uncertainties are derived based on a literature study and characterize the actual situation/ data. the sensitivity analysis is performed for different probabilistic models of the influencing parameters. to quantify the relative influence of the parameters, their first order sensitivity index is calculated. based on the computed sensitivity index of the operational expenditures related to the support structures, it can be concluded that an optimization of the structural integrity management (sim) aiming to reduce the operating costs may have negligible influence on the lcoe. a similar result has been obtained in a case study considering the effect of d-strings on the lcoe of offshore wind turbine blades [23]. however, a significant influence of a service life extension on the lcoe is found. it can thus be concluded that the structural integrity management should be directed towards a service life extension. this may be achieved by means of monitoring the support structural integrity, i.e. by optimized monitoring systems to be utilized to optimized the time and location of measures such as repairs, which ensure the reliability of the wind turbine support structures beyond the original service life. it 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https://peak-wind.com/insights/opex-benchmark-an-insight-into-operational-expenditures-of-european-offshore-wind-farms/ https://peak-wind.com/insights/opex-benchmark-an-insight-into-operational-expenditures-of-european-offshore-wind-farms/ https://peak-wind.com/insights/opex-benchmark-an-insight-into-operational-expenditures-of-european-offshore-wind-farms/ http://www.xn--drmstrre-64ad.dk/wp-content/wind/miller/windpower%20web/en/tour/wres/weibull.htm http://www.xn--drmstrre-64ad.dk/wp-content/wind/miller/windpower%20web/en/tour/wres/weibull.htm http://www.xn--drmstrre-64ad.dk/wp-content/wind/miller/windpower%20web/en/tour/wres/weibull.htm https://doi.org/10.17736/ijope.2017.il49 acta polytechnica ctu proceedings 36:90–98, 2022 1 introduction 2 levelized cost of energy 3 sensitivity analysis 4 numerical study 5 summary and concluding remarks references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0620 acta polytechnica ctu proceedings 38:620–629, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague investigating the change in preference of living and transport during covid-19 oskar fahlstedt∗, winnie ma norwegian university of science and technology, høgskoleringen 7a, 7034, trondheim, norway ∗ corresponding author: oskar.fahlstedt@ntnu.no abstract. the covid-19 pandemic meant that people had to perform most of their daily activities from home. this study intends to investigate if a period of working from home has changed living preferences concerning housing size and location. the data was collected through a web-based survey. the literature is divided regarding the potential benefits of reduced commuting due to increased home office utilization. there is a prevalent risk of rebound effects as saved travel time might lead to increased travel on leisure, which could counterweight the benefits of reduced daily commuting in cities. most of the respondents in this study want to work from home 2–3 days a week. the respondents that desire to move to a less central location state the lower square meter price as one important reason. analyses in prevailing research did not investigate the preference for larger homes combined with increased home office use. future studies should also investigate the desire for larger homes to achieve a holistic evaluation of the emissions related to telecommuting. keywords: covid-19, teleworking, telecommuting, emissions. 1. introduction the covid-19 pandemic has affected many people either directly or indirectly. in attempts to reduce the spread of the virus, a lot of people have worked from home aided by modern information and communication technology (ict). the long-term societal effects of the pandemic remains unclear, but it illuminate the importance of the the sustainable development goals (sdgs) for human and planetary prosperity [1, 2]. however, the results of impeded mobility and inactive industries have reduced greenhouse gas (ghg) emissions and improved the air quality in cities, which benefits sdg 13 (climate action) [1–3]. moreover, albeit good news, a sensible assumption is that this is a deviation whilst society will return to its old behavior patterns once the pandemic is over [3, 4]. nevertheless, it is imperative to investigate this endeavor as utilization of the home office for a long time might affect the perception people have of their work and housing situation [4, 5]. it might contribute to a situation where people demand larger homes due to a continuation of the home office. as housing prices generally decrease with greater distance to city centers, it might also increase urban sprawl. furthermore, as larger homes demand more energy and material whilst residents in the periphery, to a more considerable extent, depend on their cars, the positive environmental effects might be reversed instead. this paper intends to investigate this notion and ask the following research questions: (1.) how has a period of working from home changed people’s preference of distance to work and housing size? (2.) what potential challenges can derive from a changed preference in commuting and housing size? 1.1. literature review when investigating the potential positive and negative implications of increased utilization of the home office it is sensible to mention the aspects of “teleworking” and “telecommuting”. the literature defines a teleworker as a person that uses information and communication technology (ict) to work from another location than the physical workplace [6, 7]. furthermore, according to cerqueira et al. [8] a distinction is to be made between a teleworker and homeworker, with the former working part-time from home and part-time from work while the latter identifies the home as the physical workplace. o’keefe et al. [9] defines telecommuters as anyone that works from home at least one day of the workdays during the week. thus, a telecommuter is a person that partially works from home at least one day a week and commutes to work all other days not working from home [7]. in this paper, the term telecommuter is adopted and used in the results to describe a person that works from home at least one day of the workweek. thus, the term is overarching and includes the aspects of the term teleworker as well. moreover, the characteristics of teleworking and telecommuting are not limited to a change in environmental burdens such as emissions from commuting or energy use in buildings, there are also social aspects to consider [10]. however, the scope of this paper focuses on the potential environmental benefits or burdens of changed living and transport preferences in a situation when more people are choosing to work from home. 620 https://doi.org/10.14311/app.2022.38.0620 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 investigating the change in preference of living and transport . . . 1.1.1. positive aspects of increased use of home office the advance in ict has enabled the home-space to become an environment for work as well as any office for many people. and some potential benefits of this type of work arrangement are the reduced congestion in traffic and the need for fewer square meters in office spaces [6, 11, 12]. furthermore, with less commuting, there are direct environmental benefits like the reduction in greenhouse gases (ghg) from vehicles with internal combustion engines (ice) [13]. the indirect benefits deriving from reduced demand for office space are reduced embodied (less material needed) and operational emissions (decreased energy demand for heating and cooling) [14, 15]. an analysis of travel patterns based on mobility reports provided by google during the covid-19 pandemic combined with multi regional-input output data (mrio) of the global economy demonstrates that there is a significant reduction potential of environmental stressors. this is the conclusion in the work of pomponi et al. [16] as the developed scenario for increased teleworking together with 50 % reduction in energy demand in offices would lead to a 5 % reduction of global ghg emission. the authors put into the context of the paris agreement and state that it is half of the yearly decrease in ghg needed to align with the goal for 2050 [16]. through stakeholder involvement and backcasting, quantitative scenarios were developed and calculated in the article by godínez-zamora et al. [17]. the authors conclude that telecommuting is essential in reducing transportrelated emissions in cities. in a framework developed by shabanpour et al. [12] the consequences of teleworking are investigated in three adoption levels and the results show that daily co2 and pm25 emissions can be reduced with 696 (-0,7 %) and 167 (-1,14 %) tonnes a day on average 0,7 % and 1,14 % in their study performed in chicago. in a study performed by giovanis [18] the change in air quality in a swiss city was analyzed in a scenario where increased teleworking decreases the traffic volume. the result after analysis of demographic household data indicates that it reduces toxic pollutants as fewer commute to work and should be included in sustainable urban planning strategies. o’keefe et al. [9] performed a study based on the results of census data from the greater dublin area and complemented it with a survey for deeper analysis. they determine that 27 % of the population work from home at least once a week and that a majority of the respondents (53–74 %) prefer to work from home at least once a week. thus, the authors conclude with the use of proxy values for kg co2 emissions that a scenario with 20 % of the population working from home one day a week for a year, the total saving would amount to 60,000 tons of co2 [9]. it amounts to about an 8.6 % reduction in total co2 emissions deriving from transport in dublin [9, 19]. finally, güereca et al. [13] investigate carbon abatement strategies for a research institute on a university campus in mexico city, and the results from a processbased life cycle assessment approach demonstrate that a scenario with reduced commuting will reduce the total co2 emissions. if half of the staff and students would work remotely two out of five days, then the total emissions would be reduced by 15 %, and combining this with increased car-sharing would lead to a 22 % reduction in total emissions. 1.1.2. potential negative aspects of increased use of home office zhu and mason [15] investigate the potential telecommuting has for reducing travel distance and indirectly ghg-emissions emissions in the usa by scrutinizing data gathered from the us national household travel surveys (nhts). the results reveal that co2 emissions have increased for the telecommuting population. there might be a potential rebound effect when telecommuting in combination with increased fuel efficiency induces people to accept longer distances to work and increase non-work-related trips instead. similarly, cerqueira et al. [8] discovered that co2 emissions deriving from teleworkers are higher than non-teleworkers due to rebound effects of increased distances for commuting and non-related work trips. there is a potential for an increase in this pattern when developments in ict make the home office more efficient. furthermore, if more people utilize their home as office space the home the electricity consumption will increase and time saved from not commuting leaves people with more time to travel longer distances in their spare time [20–22]. in the findings from cerqueira et al. [8] co2 emissions from teleworkers are higher than non-teleworkers due to increased kilometers needed during the days traveling to work. it also identifies men as being responsible for a higher share of direct emissions since they traveled long distances by car. in comparison to non-telecommuters, the telecommuter has a larger travel budget which entails that they are less dependent on having access to a physical workplace and other amenities [21]. zhu et al. [21, p. 25] describe them as “footloose” as working from home enables a lifestyle that is decoupled from geographical locations and consequences might be increased instead of decreased travel demand. as aforementioned, reduced commuting enables more time for non-work trips [20–22] and alonso et al. [22] demonstrate the potential rebound effects as vehicle kilometers (vk), energy use, and emissions increase in a scenario where increased teleworking is assumed. 1.1.3. holistic approach to increase telecommuting urban planners need to be cautious of how neighborhoods are designed and be aware of the consequences that demographic aspects have for the adoption of teleworking among inhabitants. the socio-economic, neighborhood characteristics and accessibility are factors that determine where inhabitants chose to live [7]. additionally, urban forms are also influential as high street connectivity, access 621 oskar fahlstedt, winnie ma acta polytechnica ctu proceedings to public transport, and density make fewer people work from home. for neighborhoods with higher land use per person, irregular street patterns, and less accessible public transport the utilization of the home office is higher [23]. if given the opportunity to work from home at least once a week workers are less inclined to let the distance to work be a deterministic factor as workers accept increased commuting time (5 % longer) between home and the work place [6]. allen et al. [7] provides a rich review of the literature when investigating the multifaceted outcomes of telecommuting and concludes that there is not a clear answer to whether telecommuting is beneficial or not. a multidisciplinary approach is sought as the topic of telecommuting must be addressed holistically. thus, not only consider the environmental but also acknowledge the softer social aspects that influence the adoption of working from home. the demand for local amenities such as affordable and diverse housing, ict, schools, daycare, and leisure influences the decision to work from home or not [7, 10]. for example, no accessibility to kindergarten might leave people with no other option than combining child care and working from home [7]. there is also an aspect of gender since women care for children to a higher extent than men and are more likely to have to combine work with childcare [7, 10]. the economical aspects are also prevalent. for example, tax reductions for the one working from home due to infrequent use of cars compared to a regular day commuter using a car to and from work [24]. 2. method and sample a self-assisted web-based survey was used to collect the opinions and experiences of people living conditions in trondheim (norway) during the pandemic. the data collection lasted two and a half months from mid-june to the end of august 2021. the samples were created through convenience sampling by distributing one thousand flyers in residential areas (mailboxes) and public spaces. the use of a web-based method was advantageous in mid-pandemic times regarding infection risk, being spatial-temporal flexible, and in collecting a number of respondents within the project’s framework. anonymity was an important aspect of the study, as for the layman transportation research often is associated with strong intentions towards sustainable transport which may contribute to social desirability bias; skewed results in favor of what is deemed “correct” behavior. from the information on the flyer, the survey could be accessed by scanning the qr-code or typing the url address. the residential areas were of mixed housing types and were centrally located in the city. there is a propensity of participation bias as the accessibility to mailboxes varied according to housing types (difficult to access mailboxes for some apartment complexes). aside from residents, the flyers were placed in public areas like at bus stops, outside groceries stores, public bulletin boards, and other busy areas. thus, the geographical scope of the study is not limited to the residences in central areas of the city but expands outwards to respondents that live further out but that happened to have errands in the central area of the city. responses from visitors or tourists were discarded from the results. the approach used when distributing the flyers means that it is not possible to provide any response rate as multiple people might have been informed through the same flyer. the survey consisted of four parts; the first part covering demographic information of the respondent, the second subjective questions on residence location and other facilities, the third investigating the prospect of work-from-homes, while the fourth and last part collected information regarding travel habits covering the primary mode and commute time, and preferred travel time for various activities. 3. results and discussion 3.1. housing location the respondents were asked if their preferences towards residence location had changed regarding centralization or decentralization. 57 % of the respondents stated unchanged preferences, about 19 % state they want or already have moved more centralized during the pandemic, and the rest 24 % state wanting to or already moved further away from the city center. furthermore, the respondents that stated changed preferences were asked to rate the degree of influence on these preferences on a list of different aspects of travel activity, living conditions, or residence characteristics. among the respondents stating an increased preference for centralization, the accessibility to public transport and mobility by active modes was deemed important aspects, see top rows in figure 1. on the opposite range, aspects having relatively low influence are “closeness to health services” and the aspect of “less residence area to maintain”. oppositely, the respondents stating an increased preference for more decentralized residence location, favor residencerelated aspects, see figure 2. living further away from heavy traffic areas as well as access to a garden are also deemed important aspects. on access to parking space for the private car, only 47 % deemed it as influential. the increased rate of telecommuting has not induced the strongest effect on preferences, having 60 % stating an effect, where only 25 % percentage being “substantial effect”. food delivery services and digital alternatives for socializing were not assessed as very influential aspects regarding location preferences. the aspect of residents wanting to move to larger homes was not covered in the literature as the consequences of telecommuting and teleworking were foremost discussed in terms of emission from reduced or increased travel [8, 9, 15, 16, 18, 21]. the potential increase in energy use in the homes was discussed [20–22] but the demand for larger homes due 622 vol. 38/2022 investigating the change in preference of living and transport . . . figure 1. the assessed effect of each aspect on preferences for centralization (n = 63). figure 2. the assessed effect of each aspect on preferences for decentralization (n = 80). to telecommuting was not covered is perplexing since it is an important aspect when discussing emissions in the built environment. however, it is imperative to investigate as holistic assessments are sought for when discussing telecommuting [7]. the impact of house size adds another dimension to the debate of the impacts of telecommuting. this is since residents living in larger homes felt more content with their home office situation. regarding telecommuting, the respondents preferring centralization have on average fewer commuting days and the least willingness or possibility to telecommute post-pandemic compared to the “unchanged” and “less central” respondents. these results may indicate a tendency of more positive attitudes towards telecommuting when having preferences towards decentralization rather than centralization or vice versa, an attitude on telecommuting that depends on location preferences. in general, 48 % of the respondents stated being just as content with their residence during the survey response moment (mid-pandemic) compared to prepandemic times. about 28 % stated being less content and 20 % being more content, while the rest stated not being sure. in combination with other variables, level of contentment seem to have no clear correlation with stated residence density, residence type, residence size, age, household type, gender, and income. 3.2. transport mode and travel times the respondents were asked to assess the maximum acceptable travel time for various trip purposes independent of their current residence location. not all purposes were applicable and the least relevant were kindergarten, primary school, and high school by only 30 % of respondents, see figure 3. the lowest average accepted travel time is for bus stops at 6 minutes, whereas the next lowest is grocery stores at 9 minutes. on the opposite end, trips to the workplace or cinema/theaters have on average an accepted travel time of 25 minutes. the fact that workers are willing to accept longer commute times makes an increased adoption of home office an interesting scenario in the post-pandemic future. the respondent’s general priorities in residence location may be based on their assessment of important aspects and their maximum acceptable travel time for different destinations. it 623 oskar fahlstedt, winnie ma acta polytechnica ctu proceedings figure 3. the average maximum preferred travel time by the various trip purposes. the dots’ sizes represent the share of applicability for the respondents. reveals that commute times may be one of the more flexible variables compared to trip purposes with much lower acceptable travel times. the lower the travel times, the more restrictive the vicinity to this destination becomes in residence location compared to other activities/destinations given the assumption that each destination is equally important for the individual. it is important to notice that the time assessment states a desire or a preferred estimate, but do not tell us the relative importance of each trip purpose when facing the dilemma of fulfilling the max criteria of several different destinations (e.g. more than 6 minutes to bus stop versus more than 14 minutes to the gym). to assess the importance of the destinations/activities, one could conduct a stated preference survey followed by a calculation of the travel time elasticizes per activity/destination. while not accounting for the individual destination importance, our sample reveals public transportation to be the dominant primary mode for both work and leisure purposes, making the proximity to bus stops at 6 minutes quite relevant as confirmed in the assessment of aspects that affect the preference for centralization. thus, the potential impact of increased car use due to teleworking [8, 15, 21, 22] is not revealed in this study as a geographical decoupling is potentially not viable when respondents are dependent on public transport services as much as they were in this study. the requirement for bus stops nearby reveals that a relocation further away is not preferable as the frequency of bus stops usually decreases with a lower density of residencies. the average commute is 23 minutes while the average accepted commute is 27 minutes, see figure 4 for the distribution of the time difference between the respondents’ acceptable commute times and their actual commute times. the density figure 4. the density distribution of the time difference between maximum acceptable commute times and actual commute times (n = 182). figure 5. transport mode distribution each for work and leisure purpose. distribution shows a peak at 0 minutes, indicating that a majority of respondents do not prefer any additional travel time by any means. the graph has more volume on the right side which reveals that most of the workers have commute times within their maximum travel time. though, it is also apparent that there are some incidents of negative values (the other side), indicating that a few individuals are currently commuting longer than they want to. the distribution of primary modes for commute trips (n = 183) show that public transport dominate with 45 %, see figure 5. mind the newly emerging mode e-scooters make up almost 2 % of the main modes in respect to commuting. given car as the main mode, the total pool of cars accessible (can be more than one car) consist of 40 % electrical powered cars, 36 % gasoline, 20 % diesel, and the rest powered by alternative energy sources. on a national level, 12 % of the private cars are electric-powered, but this share is higher for the average car commuter living in urban areas that cross a toll ring – as an electric car pays no fees passing the toll per this date 2022 [25]. the mode distributions are similar across genders, households, and education levels. the mode distributions are similar across genders, households, and 624 vol. 38/2022 investigating the change in preference of living and transport . . . education levels. tendencies for a higher share of car use are found among households with children, households with higher income, single-family residences, or row houses compared to apartments. the modal distribution seems to be strongly related to the travel times for work trips, with active modes as the dominant mode choice for shorter trips until public transport takes over. the car share does not vary the same as the other modes, where a possible explanation for this is its ability for trip chains (e.g., escort trips to school/activity), having space to bring more people and goods, reducing any walking and waiting time and requiring marginal physical effort. for leisure trips (n = 338) the mode distribution was 40 % for public transport, still a great share, but less than the modal distribution for work trips, see figure 5. given car as the main mode, 37 % of the accessible cars are powered by gasoline, 37 % by diesel, 20 % electric, and 7 % alternative energy sources. the percentage of car use is higher for leisure than commutes, which is a bit surprising as the fact that students’ mode choices are included in leisure (and not in the distribution for commute trips) was thought to have a reducing effect on the share of car usage. it was revealed that the cause of this was mostly due to workers having public transport as their main mode for commutes but switching to the car as their leisurely main mode. this is interesting as some findings in the literature indicated that car trip distances increase when working from home [8, 22]. when asked of walking time to the closest public transport stop/station, 76 % state less than 5 minutes and 20 % from 5 to 10 minutes. there was no clear correlation between mode choice and vicinity to public transport stops. 3.3. working from home and their willingness to move when asking full-time, part-time workers and students (total n = 302) about their preferences regarding telecommuting, 50 % were positive to telecommuting (n = 148), though of those 7 % and 6 % stated a wish for telecommuting but experienced challenges with inadequate workspace at home and not being sure of the employer would allow it, respectively. 30 % stated not wanting to telecommute post-pandemic, and 5 % was not sure. out of those wanting to work from home, 12 % stated a wish of working max 1 day a week, 65 % was willing to work up until 2–3 days a week and 18 % stated a willingness for 4–7 days a week. there is some inconsistency among the answers, as 4.3 % answered a willingness of 0 days a week after stating a willingness for telecommuting. zhu et al. [21, p. 25] described teleworkers as “footloose” since they are less dependent on the physical location of the workplace and amongst the respondents, half of them state a wish to, some extent, continue to work from home after the pandemic. the results show that among the full-time workers there is a greater share willing to telecomfigure 6. willingness to move for increased residence size given the possibility to telecommute in respect to number of days telecommuting mid-pandemic. mute compared to part-time workers. the part-time workers state more often that their work tasks cannot be conducted from home. of those n = 163 willing to telecommute post-pandemic (50 %), aside from asking about the willingness of the number of days telecommuting, they were asked about their willingness to move for increased residence size and if they’re willing to move further away from their current workplace. when asked about their willingness to move for a bigger residence given the possibility of telecommuting, 59 % were willing, 29 % were not, and 12 % were not sure. 41 % (n = 64) were willing to move further away from their workplace, 40 % were not, and 19 % were not sure. it is as expected that some would be willing to increase the residence size without the compromise of moving further away, though our results also show the opposite, where some individuals want to move further away without increasing the residence size (n = 8). the results show different relations between location preferences and days telecommuting compared to “willingness to increase residence size” and days telecommuting, as a viable increase in residence often involves relocation less central (not necessarily, but often a correlation). given the possibility of teleworking in the future, respondents answering “no” to a willingness of increased residence had an overall higher median number of days telecommuting during the pandemic compared to those positive to increase residence size, see figure 6. the results on the hypothetical post-pandemic telecommuters indicate that their willingness to move further from their workplace is not influenced by their work-from-home frequency mid-pandemic, as those willing to move further and those not wanting to move have the same wishes of working from home in the future. there are as many reasons to move residence as there are individuals, but the contentment level of their residence may be an overarching factor for a general indication of their willingness to move. this study asks respondents about their willingness to move residence relatively to their workplace, and further questioning on transport are only asked those who report a willingness to move further away from their workplace as opposed to those who do not want that. 625 oskar fahlstedt, winnie ma acta polytechnica ctu proceedings this is because moving further away would generally increase commute times, reduce the propensity for active modes of public transport and therefore increase traffic-related burdens (e.g. rush hour, congestion, local pollution). results indicate that the respondents being less content with their residence have a higher willingness to move given the possibility of telecommuting. though, the effect of contentment is not unidirectional as those being “more content” show a level of willingness lying just in between the “less content” and “unchanged”, when it would have been expected to show the lowest willingness (below the level of those answering “unchanged”). 3.3.1. transport and travel time given telecommuting the respondents that answered a willingness to move further away from their workplace given the possibility of telecommuting, were asked to assess the potential acceptable increase in commute times. for all transport modes except biking, there was stated a range of increased travel time, see figure 7. it is noticeable how some pedestrians would be willing to walk that much longer if not thinking of changing mode in such a context. mind that each sample is limited and a few extreme values may therefore have a relatively big influence on the aggregated results. a look at the current commute times and the willingness of increasing these show no clear tendency between the two variables. the number of stated maximum days of telecommuting post-pandemic do not seem to have a clear relation with the willingness for longer commute times as showing big variations among the answers. the group of workers (n = 64) stating a willingness to move further from their workplace (the 41 % positive to telecommuting), though a small group, have a transport distribution that reveals, compared to the distribution for all workers, that pt and e-bike users were slightly over-represented, pedal bike users underrepresented, pedestrians and car users were “balanced” represented in their willingness to move further away from their workplace. this is a bit surprising as one would expect the ones having the least distancesensitive modes (e.g. car) to show a greater willingness to move decentralized. given this study’s sample, among students and workers, 4–6 % self-predict increased car use given a reduction in pt/ biking (when willing to move further away from the workplace due to telecommuting). with a bigger sample size, it would be possible to look closer into which primary modes that predict the biggest modal shift. the potential benefits of telecommuting are foremost reported in the form of reducing emissions related commuting transport and traffic congestion. literature has shown that telecommuting reduces commute frequency, but may have an increasing effect on leisure trips. this since time saved on commuting from might increase car use instead as it is a popular figure 7. willingness for longer commute times given new residence location due to telecommuting, grouped by stated primary transport mode. mode of transport for leisure [8, 15, 21, 22]. although, the literature describes how the car is utilized to a further extent for spare-time trips than commuting there was no clear distinction of this rebound effect in the results from the questionnaire as a majority stated public transport and walking as the primary mode for leisure. a possible effect of fewer commute trips by pt users is the “lower value for money” when traveling on the same period ticket for fewer trips – which may result in omitting the purchases of period tickets altogether in favor of private modes like car or bike. other studies highlight that telecommuters tend to commute further when going to their workplace, indicating that their homes are located further from work compared to non-telecommuters which is problematic with the primary mode being car [8]. the respondents that used motorized transport stated that them working more from home would make them accept longer commutes. about half of the workers were willing to work from home 3–5 days a week post-pandemic and the accepted increase in commute times was 35– 45 minutes (depending on transport mode) which is significantly higher than suggested in literature [6]. however, the sample size is limited and has volatile results due to the marginal influence of a few extreme values. 3.4. limitations the study was conducted in a university city resulting in a data collection recruiting an overweight of students. the survey sample, although recruited as evenly as possible within limited time and financial resources, cannot be regarded as demographically representative of the entire population, but rather as indicative of the residential experience in a city during the pandemic. as a limitation of this study and discussion of the article, the presented analyses are mostly overarching, and rather point into general directions or tendencies of main mechanisms. as such, there is a demand for future research to go beyond the descriptive results presented in this article and focus on finding potential variables with the least or most influence on the observed tendencies. 626 vol. 38/2022 investigating the change in preference of living and transport . . . 4. conclusions the conclusion sets out to answer the two research questions asked in the introduction of this study. 4.1. how has a period of working from home changed people’s preference of distance to work and housing size? in total there are almost equal shares of respondents having changed preferences towards centralized and decentralized locations, where people favoring decentralized locations have had and wish to continue telecommuting more than those with other preferences. given the 50 % respondents being positive towards future telecommuting, the average willingness of maximum days telecommuting is 2–3 days a week. about three out of five future telecommuters state a willingness to increase their residence size due to telecommuting. there is no clear correlation between the frequency of home office mid-pandemic and the wish for a larger residence, though it is worth mentioning that resourceful individuals more often possess the types of jobs more suitable for home office; one of several potential mechanisms. but for the number of days wanting to telecommute post-pandemic, respondents willing to move had a higher median day per week of maximum days telecommuting than the individuals who did not want to move. 4.2. what potential challenges can derive from a changed preference? the literature is divided regarding the environmental benefits of telecommuting. an increase in telecommuters reduces the emissions related to omitted commute trips, but a possible rebound effect may occur in the form of longer leisure trips. in total, the changes in travel habits may cancel each other out or the rebound effect outweighs the benefits. the potential emissions from the changed preferences amongst the respondents were not addressed in this paper but scenario testing with a focus on embodied and operational emissions from transport and housing should be investigated further. the findings from the study demonstrated that respondents preferred to travel by public transport for commuting and leisure purposes. thus, infrastructure and transport services are important factors even if telecommuting increases. the rebound effect discovered in the literature was not prevalent in this study as public transport is used more than cars. the desire for larger homes when working from home adds another dimension to the potential rebound effects of the home office. the benefit of the reduction in required office space should be analyzed with the desire of people wanting to increase the size of their homes. finally, some suggestions for further research: (1.) a larger cross-sectional study similar to this study in cities with other work cultures; (2.) scenario testing with a focus on embodied and direct emissions from groups with various work arrangements; (3.) the potential consequences of the desire for larger homes when working from home should be investigated further. 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https://www.toi.no/getfile.php?mmfileid=53659 vol. 38/2022 investigating the change in preference of living and transport . . . a. appendices figure 8. share of respondents having telecommuting before, during and wanting to after the pandemic. 629 acta polytechnica ctu proceedings 38:620–629, 2022 1 introduction 1.1 literature review 1.1.1 positive aspects of increased use of home office 1.1.2 potential negative aspects of increased use of home office 1.1.3 holistic approach 2 method and sample 3 results and discussion 3.1 housing location 3.2 transport mode and travel times 3.3 working from home and their willingness to move 3.3.1 transport and travel time given telecommuting 3.4 limitations 4 conclusions 4.1 how has a period of working from home changed people's preference of distance to work and housing size? 4.2 what potential challenges can derive from a changed preference? references a appendices doi:10.14311/app.2021.30.0063 acta polytechnica ctu proceedings 30:63–68, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app influence of earth composition on adsorption capacity of methylene blue dye tereza plaček otcovská∗, barbora mužíková, pavel padevět czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: tereza.otcovska@fsv.cvut.cz abstract. final properties of unfired earth are influenced by composition of earth mixtures. methylene blue test could be useful method for analysis of earth composition. this would facilitate of designing unfired earth building structures because composition of natural earth is various. principle of methylene blue test is measure of amount of adsorbed methylene blue dye by clay. the essential component of earth mixtures is clay because clay fulfils a function of binder. experimental measure of adsorption capacity of individual kind of clay is described in this paper. adsorption capacity was investigated in montmorillonite, kaolinite, illite-kaolinite and illite clays. the obtained results show that the adsorption capacity of clays is significantly different. keywords: adsorption capacity, clay, methylene blue dye, methylene blue test. 1. introduction unfired earth is a building material which had become minor in 2nd half of 20th century because of modern building materials such as steel or concrete have have become readily available. nowadays, the unfired earth is getting back to forefront of interest for its properties that fit into principles of a sustainable building and improving microclimate in interiors [1– 4]. earth mixture is a basic material for production of unfired earth constructions. uncertain composition of natural earth is the main reason why using of unfired earth in construction is complicated. it is because of final properties of unfired earth constructions are greatly influenced by the earth composition [4, 5]. methylene blue test is described in standard čsn en 933-9. methylene blue test could be a good method for analysis of the earth composition. basic idea of this research is that every kind of clay (main component of earth mixture) has different adsorption capacity of methylene blue dye. if adsorption capacity of individual clay is known, this test could be a good method for analysing earth composition [5–8]. this paper is focused on experimental determination of adsorption capacity of clay by methylene blue test. four kinds of clay were investigated (montmorillonite, kaolinite, illite-kaolinite, illite). minimum number of measurements for each investigated material was 3. the result of this research is the determination of the adsorption capacity of mentioned clay and determination of dependence between the amount of adsorbed dye and amount of clay in earth. 2. description of experimental measurement a burette, a beaker, a filter paper, a magnetic stirrer and a glass rod are basic equipment for methylene blue test. solution of methylene blue dye (methylthioninium chloride c16h18cln3s) by concentration 10 g/l is used. earth samples are produced from earth and distilled water. methylene blue solution is inserted into earth sample by a burette (fig. 1). figure 1. an earth sample with methylene blue solution. earth sample with the solution is stirred on a magnetic stirrer 1 minute for each 1 millilitre of solution. after that a drop of the mixture is placed on a filter paper. this process is repeated and each sample drop with increasing amount of methylene blue solution is marked (fig. 2). methylene blue test is finished when a blue ring spreads around a drop sample (fig. 3) because clay has already adsorbed the maximum amount of methylene blue dye. amount of dye adsorbed depends on a kind and amount of clay. amount of dye adsorbed is determined by the eq. 1. amount of dye adsorbed is defined as the weight amount of dye in grams that is adsorbed by 1 kilogram of test material. 63 http://dx.doi.org/10.14311/app.2021.30.0063 http://ojs.cvut.cz/ojs/index.php/app t. plaček otcovská, b. mužíková, p. padevět acta polytechnica ctu proceedings figure 2. a drop of the mixture is placed on a filter paper. figure 3. a result of methylene blue test. mm b = ms.mm b,1ml.1000 mes (1) 3. investigated material earth samples are produced in laboratory from sand, clay powder (fig. 5) and distilled water. a grain curve of sand is shown in fig. 4. 4 kinds of clay are tested (montmorillonite, kaolinite, illite-kaolinite, illite). industrial marking which is used in this paper too and chemical composition are shown in table 1. earth samples are suspension of clay and sand in distilled water. clay and sand are mixed in prescribed ratio (table 2) and it is solid part of earth figure 4. grain curve of sand. figure 5. clay powder. samples. this solid part is always 15 g. distilled water is always 75 g. distilled water serves as a medium for adsorption. 4 clay/sand ratios are used for each kind of clay. 54 measurements are performed on 16 types of samples. the minimum number of measurements for each sample type was 3. table 2 shows composition of samples, amount of components and a number of measurements. mark of a sample contains mark of clay and percentage number of clay in a sample. table 2 shows mark of samples. 4. adsorption capacity of clay adsorption capacity is very different for each kind of clay. the highest adsorption capacity is found for montmorillonite clay (gem), lower for kaolinite clay (b4), even lower for illite-kaolinite clay (kr) and the lowest for illite clay (agl). table 3 and figure 6 shown results from methylene blue test. maximum average amount of adsorbed dye 88 g/kg (σ = 3.8 g/kg) is calculated for gem-100. the clay 64 vol. 30/2021 earth composition and mbd adsorption capacity kind of mark of amm sio2 al2o3 fe2o3 tio2 cao mgo na2o k2o clay clay [ml/g] [%] [%] [%] [%] [%] [%] [%] [%] montmorillonite gem 17 50.51 31.2 3.37 0.86 0.4 0.42 0.08 1.62 kaolinite b4 4.5 58.13 24.29 4.7 1.21 0.13 0.5 0.5 3.96 illite-kaolinite kr 8.4 59.31 24.71 3.37 1.09 0.19 0.4 0.3 2.82 illite agl 29.4 56.57 18.4 9.72 1.16 1.12 2.54 0.18 2.91 table 1. chemical composition of used clays. clay mark of sample clay/sand ratio amount of distilled water amount of clay amount of sand number of measurements [−] [g] [g] [g] [−] montmorillonite gem-100 100/0 75 15 0 3 gem-75 75/25 75 11.25 3.75 4 gem-50 50/50 75 7.5 7.5 3 gem-25 25/75 75 3.75 11.25 5 kaolinite b4-100 100/0 75 15 0 3 b4-75 75/25 75 11.25 3.75 3 b4-50 50/50 75 7.5 7.5 3 b4-25 25/75 75 3.75 11.25 3 illitekr-100 100/0 75 15 0 3 kaolinite kr-75 75/25 75 11.25 3.75 3 kr-50 50/50 75 7.5 7.5 4 kr-25 25/75 75 3.75 11.25 5 illite agl-100 100/0 75 15 0 3 agl-75 75/25 75 11.25 3.75 3 agl-50 50/50 75 7.5 7.5 3 agl-25 25/75 75 3.75 11.25 3 table 2. clays used for experimental measurements, their designation and composition. figure 6. amount of adsorbed methylene blue dye by clays. contained in a sample is montmorillonite and clay is 100 % of the solid part of a sample. on the contrary, the lower value of adsorbed dye 4.6 g/kg (σ = 0 g/kg) is determined for agl-25 sample. the clay contained in agl-25 sample is illite and clay is 25 % of the solid part of a sample. compared to montmorillonite clay, the adsorption capacity of kaolinitic clay is 34.3 % lower (e.i. 65.7 %), illite-kaolinite is 63.5 % lower(e.i. 36.5), illite is 78.2 % lower (e.i. 21.8 %). figure 6 shows that the same trend was also observed for the samples with a clay content of 75 %, 50 % and 25 %. the percentage difference in adsorption capacity for all clays is given in table 4. obtained values of adsorbed dye by individual samples are interpolated by a linear regression curve. it can be seen from figure 7 that dependence of adsorption on amount of clay is linear for all kinds of clays. 4.1. confidence intervals from the obtained adsorption values, limits of the confidence interval l1 and l2 were calculated by eq. 2 (α = 0.05). calculated confidence interval limits are given in table 5. l1,2 = x ± σ · tk√ n (2) confidence interval is calculated for all measurement values of adsorption dye. figure 8 shows dependence between amount of adsorbed dye and amount 65 t. plaček otcovská, b. mužíková, p. padevět acta polytechnica ctu proceedings clay mark of sample average amount of adsorbed dye standard deviation relative standard deviation [g/kg] [g/kg] [%] montmorillonite gem-100 88.0 3.8 4.3 gem-75 65.3 5.6 8.6 gem-50 41.8 6.9 16.4 gem-25 17.3 3.5 20.1 kaolinite b4-100 57.9 2.0 3.5 b4-75 44.5 1.7 3.7 b4-50 29.3 3.4 11.6 b4-25 14.5 0.7 4.5 illitekr-100 32.1 2.1 6.6 kaolinite kr-75 25.5 3.6 14.2 kr-50 18.0 1.4 7.6 kr-25 8.7 1.1 12.4 illite agl-100 19.2 0.6 3.3 agl-75 13.9 1.3 9.5 agl-50 9.5 1.0 10.7 agl-25 4.6 0.0 0.0 table 3. results obtained by methylene blue test. amount of adsorbed dye gem b4 kr agl amount of clay [g/kg] [g/kg] [g/kg] [g/kg] in a sample [%] [%] [%] [%] 15.00 g / 100 % 88.0 57.9 32.1 19.2100 65.7 36.5 21.8 11.25 g / 75 % 65.3 44.5 25.5 13.9100 68.1 39.1 21.2 7.50 g / 50 % 41.8 29.3 18.0 9.5100 70.0 43.0 22.6 3.75 g / 25 % 17.3 14.5 8.7 4.6100 84.0 50.4 26.7 table 4. comparison of the amount of adsorbed dye. of clay in sample with confidence interval area for each clay. the curves of adsorption are created by combining the calculated values of adsorption for amount of clay in samples 25 %, 50 %, 75 % and 100 %. extreme values of confidence interval area are join between the calculated values of the confidence interval limits l1, l2. it can be seen from figure 8, that illitic clay (agl) and illitic-kaolinitic clay (kr) can be well distinguished from each other in an amount of clay in solid part at least 25 %. confidence intervals of montmorillonite (gem) and kaolinite (b4) clay overlap when clay content is 25-50 %. based on figure 8, it can be assumed that the methylene blue test is applicable for amounts of 75 % and higher in the case of montmorillonite and kaolinitic clays. 5. conclusions methylene blue test is applicable to diagnose composition of earth. each of the investigated clays (montmorillonite, kaolinite, illite-kaolinite, illite) adsorbs a different amount of methylene blue dye. montmorillonite is characterized by the largest adsorption capacity. the adsorption capacity of methylene blue dye decreases for clays in the order: kaolinite, illitekaolinite and illite (fig. 6). furthermore, it was experimentally determined that the amount of adsorbed dye is directly dependent on the amount of clay in earth (fig. 7). using confidence intervals, it was determined that illitic and illitic-kaolinitic clays can be determined by methylene blue test in earth in quantities from 25 %. montmorillonite and kaolinite clay can be determined by methylene blue test in earth in amounts from 75 % (fig. 8). 66 vol. 30/2021 earth composition and mbd adsorption capacity figure 7. amount of adsorbed methylene blue dye by clay. figure 8. confidence intervals of clays. 67 t. plaček otcovská, b. mužíková, p. padevět acta polytechnica ctu proceedings clay mark of sample confidence interval [g/kg] l1 l2 g e m gem-25 13.31 21.29 gem-50 29.18 54.43 gem-75 57.54 73.16 gem-100 81.01 95.01 b 4 b4-25 13.31 15.73 b4-50 23.04 35.49 b4-75 41.45 47.53 b4-100 54.16 61.57 k r kr-25 7.49 9.96 kr-50 16.10 19.88 kr-75 18.84 32.20 kr-100 28.23 36.02 a g l agl-25 4.62 4.62 agl-50 7.61 11.31 agl-75 11.44 16.29 agl-100 18.01 20.32 table 5. calculated confidence interval limits of adsorption for clays (α = 0.05). list of symbols m b methylene blue mm b amount of mb dye absorbed [g/kg] ms amount of methylene blue solution [g] mm b,1ml amount of methylene blue in 1 g of methylene blue solution [g/g] mes amount of tested material [g] l1,2 limits of the confidence interval [g/kg] x arithmetic mean [g/kg] σ standard deviation [g/kg] tk student’s t critical values n number of measurement acknowledgements the financial support of this experiment by the czech science foundation (gačr project no. 18-10884s) and faculty of civil engineering, czech technical university in prague (sgs project no. sgs19/148/ohk1/3t/11) is gratefully acknowledged. references [1] p. jaquin, c. augarde. earth building : history, science and conservation. report, bracknell : ihs bre press, 2012. [2] agenda 21 united nations environment programme (unep). http://www.unep.org. [3] čsn. čsn 1168-1939 podmínky pro zednické a přidružené práce pozemních staveb, 1951. [4] g. minke. building with earth. ökobuch verlag, staufen, 2006. [5] i. žabičková. hliněné stavby. era 21, brno, 2002. [6] p. walker. rammed earth: design and construction guidelines. ihs bre press, watford, 2010. [7] z. weiss. jílové minerály : jejich nanostruktura a využití. karolinum,, vyd. 1. edn., 2005. [8] evropský výbor pro normalizaci. čsn en 933-9, tests for geometrical properties of aggregates – part 9: assessment of fines – methylene blue test 2013. 68 https://doi.org/10.14311/app.2022.33.0027 acta polytechnica ctu proceedings 33:27–31, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague environmental impacts of using desalinated water in concrete production in areas affected by freshwater scarcity alessandro arrigonia, ∗, tamar ophera, sabrina spatarib, valeria arosioc, heather l. macleana, daman k. panesara, giovanni dotellic a the university of toronto, faculty of engineering, department of civil and mineral engineering, 35 st. george street, toronto, ontario, m5s 1a4, canada b technion israel institute of technology, faculty of civil & environmental engineering, department of structural engineering and construction management, haifa, 32000, israel c politecnico di milano, school of industrial and information engineering, department of chemistry, materials and chemical engineering "g.natta", piazza l. da vinci 32, 20133 milano, italy ∗ corresponding author: a.arrigonim@gmail.com abstract. up to 500 litres of water may be consumed at the batching plant per cubic meter of ready mix concrete, if water for washing mixing trucks and equipment is included. demand for concrete is growing almost everywhere, regardless of local availability of freshwater. the use of freshwater for concrete production exacerbates stress on natural water resources. in water-stressed coastal countries such as israel, desalinated seawater (dsw) is often used in the production of concrete. however, the environmental impacts of this practice have not yet been assessed. in this study the effect of using dsw on the water and carbon footprints of concrete was investigated using life cycle assessment. water footprint results highlight the benefits of using dsw rather than freshwater to produce concrete in israel. in contrast, because desalination is an energy intensive process, using dsw increases the greenhouse gas intensity of concrete. nevertheless, this increase (0.27 kg co2e/m3 concrete) is small, if compared to the life cycle greenhouse gas emissions of concrete. our results show that using untreated seawater in the mix (transported by truck from the coast) in place of dsw, would be beneficial in terms of water and carbon footprints if the batching plant were located less than 13 km from the withdrawal point. however, use of untreated seawater increases steel reinforcement corrosion, resulting in loss of structural integrity of the reinforced concrete composite. sustainability of replacing steel with non-corrosive materials should be explored as a way to reduce both water and carbon footprints of concrete. keywords: carbon footprint, desalinated water, seawater, water footprint. 1. introduction water scarcity affects two-thirds of the global population [1] and climate change, economic development and population growth might worsen this issue in the near future [2]. industrial development plays an important role in the growing demand for water [2], with concrete being a significant contributor. the life cycle of concrete requires large amounts of water, from the production of the raw materials (i.e., cement and aggregates), through mixing, curing, and hydration, to equipment cleaning processes [3]. in 2012, the estimated global water consumption for concrete production was 16.6 km3 and it is expected to grow with the increase in production, by more than 40% by 2050 if no action is taken [4]. aggregate production is the main water-consuming activity in the life cycle of concrete [5], but water consumption at the batching plant is very significant. considering ready-mix concrete, up to 500 litres of water per cubic meter of concrete may be consumed, including water for washing the mixing trucks, on-site equipment and floor [6]. water used at the batching plant is typically sourced from a potable supply network, on-site wells, or harvested stormwater [6]. potable water is used to avoid any adverse effects on the properties of the concrete, such as hydration, strength development and durability performance [7]. however, using available freshwater for concrete production will exacerbate the stress on natural water resources. in water-stressed coastal countries, such as israel, desalinated seawater (dsw) is increasingly used to meet the potable water demand [8]. although reducing the stress on freshwater resources, desalination comes with an environmental cost. in qatar, for instance, 4.66 million tons of co2 were estimated to be emitted by desalination plants in 2014 [9]. one appealing option to reduce the stress on available freshwater resources, and at 27 https://doi.org/10.14311/app.2022.33.0027 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en a. arrigoni, t. opher, s. spatari et al. acta polytechnica ctu proceedings figure 1. flow chart for the product system analyzed. the different paths for the two scenarios are highlighted: desalinated and seawater. the unit processes valid for both processes (e.g., seawater withdrawal, concrete mixing, etc.) are considered unaffected by the type of water used in the mix and therefore excluded from the comparative lca. the same time avoid the impacts associated with desalination, is to replace potable water with untreated seawater in the manufacture of concrete [10]. however, the use of seawater is currently forbidden in reinforced concrete standards because it contributes to the early corrosion of carbon steel reinforcement, which is the present common practice in construction [11]. nevertheless, in recent studies, the use of seawater combined with non-corrosive reinforcing materials, such as glass fibre reinforced polymers, has shown promising results in terms of mechanical properties and durability [12]. the focus of this study is to estimate the consequences of using seawater for concrete production in israel, using a consequential life cycle assessment (lca) methodology. the results are intended to pave the way for further investigation of the feasibility and sustainability of using seawater and marine aggregates for concrete production. 2. goal the objective of this analysis is to assess the potential water and carbon footprint savings for israel if seawater was used in place of municipal grid water in concrete production. the study is aimed to set the stage for further analyses of the feasibility and sustainability of using alternative concrete designs with the ultimate goal of reducing the environmental impacts of the most prevalent construction material in the israeli construction sector. 3. scope and methods the functional unit of the study is the production of 1 m3 of generic non-reinforced concrete (100% ordinary portland cement, 0.85 water-to-cement ratio) in israel. two alternatives differing in the type of water used in the mix are compared in the analysis: municipal grid water (current practice) and untreated seawater. the market demand for concrete is assumed to be constant (18 mm3 per year [13]) and its mix design (i.e., the amounts and types of binders, aggregates and admixtures) and mechanical properties are assumed to be identical in both cases. the production process at the batching plant in this study is the only life cycle phase that differs between the compared scenarios. we use a consequential approach in this lca. unlike attributional lca, which aims to capture all environmental flows associated with the life cycle of the activity, a consequential assessment accounts only for the flows affected by a certain change in the system [14]. according to the principles of consequential lca, we consider the marginal water supply, which is the one that is most likely to react to a change in market demand. in the case of the israeli water market this is the supply of dsw, coming from five major desalination plants along the mediterranean coast. in the case of electricity, the marginal supplier in israel is under most circumstances one of several natural gas power plants [15]. we consider only the variations in impacts resulting from avoiding the desalination 28 vol. 33/2022 concrete made of desalinated water scenario unit process water footprint(l-eq./m3 concrete) carbon footprint (kg co2-eq./m3 concrete) desalinated desalination 18 0.27 distribution 4.8 0.071 mixinga 0.0 excluded total 23 0.34 seawater transportation 1.8 · x 0.016 · x mixinga 0.0 excluded total 1.8 · x 0.016 · x a the energy intensity of the mixing process, and therefore the greenhouse gas emissions and water footprint associated with the process, is assumed to be independent from the water used (dsw or sw) and therefore excluded from the analysis. since the water used for mixing was originally seawater in both scenarios, its water footprint is zero. table 1. water and carbon footprint results for the activities considered in the two scenarios (i.e., activities affected by the change in the system). variable x represents the distance (km) travelled by the truck from the desalination plant to the batching plant. process by replacing dsw with untreated seawater throughout the life cycle of concrete. the ecoinvent consequential database was used for modelling background processes in the system. the lca is limited to two environmental impact categories: climate change and water consumption. inputs and outputs to the system are considered from cradle to grave. 4. product system two scenarios of concrete production in batching plants in israel are modelled. in the base case scenario, representing current practice, water for mixing concrete is sourced from the israeli water grid. dsw is the marginal source of water in israel [16], meaning that a change in potable water demand will affect the extent of its production [17]. therefore, in the base case scenario (named desalinated scenario from now on) seawater is drawn from the sea, desalinated and distributed to the concrete batching plants via the national supply grid. in the alternative (seawater) scenario, once drawn, seawater is transported via trucks to the batching plants. if using untreated seawater proves to be a viable and sustainable solution, pipelines could be installed to deliver seawater to the batching plants. existing desalination plants inlets are assumed to be used as a means for pumping water from the sea into the trucks. the product system analysed is illustrated in figure 1. the mixing process, the delivery of concrete to the construction site, and the properties of the concrete are assumed to be unaffected by the type of water used. in future work the effects on concrete properties and on the service life of the mixing trucks (which might be reduced due to potential corrosivity of the seawater mix) will be included. 5. life cycle inventory both primary and secondary data are used for the analysis: primary data for the inputs and outputs of the israeli desalination plants and the distribution process [18]; secondary data from the ecoinvent 3.6 database [19] for the average amount of water used in the concrete mix (i.e., 170 l/m3), and for the emissions and indirect water consumption of electricity production and road transportation (e.g., water consumed in fuel production). water for washing the trucks is assumed to be reused in the concrete mix itself, as is the current common practice [13]; therefore, no additional benefits are considered for substituting municipal grid water with seawater for this purpose. moreover, additional water required at the batching plant for washing the site and for personnel needs is assumed to be municipal grid water, in both scenarios, and is therefore excluded. a threshold is calculated for the maximum distance from the seawater extraction site for which using seawater shipped by truck to the batching plant would still be beneficial in terms of carbon and water footprint. energy consumed by the desalination plants is supplied through the electricity grid, which uses natural gas combined cycle power plants (60% efficiency) [15]. 6. results and discussion table 1 presents the water and carbon footprints of the activities affected by the change in the system. results are expressed per cubic meter of fresh concrete produced at the batching plant. in the seawater scenario, impacts depend on the distance travelled by the trucks (represented by x in table 1) to transport seawater to the batching plant. for instance, the transport of 170 l of seawater by truck for the production of 1 m3 of concrete is responsible for the consumption of 1.8 l-eq. of water per km. 29 a. arrigoni, t. opher, s. spatari et al. acta polytechnica ctu proceedings the key observations from the results are: • from a consequential perspective, the water footprint of concrete mixing in israel, taking into account direct water use at the batching plant only, is considerably lower than those reported for other regions (e.g., up to 16,000 l-eq./m3 concrete in sicily, italy [6], or 185 l-eq./m3 concrete in sweden [20]). the reason is that the marginal water supply in israel is dsw, which has a water footprint characterization factor of zero since it does not negatively affect the availability of freshwater in the watersheds [21]. therefore, the only freshwater consumed in the process is the water consumed indirectly for electricity generation and fuel production. this applies to both scenarios. • the carbon footprint of desalinating seawater and distributing it to the batching plant is trivial compared with the life cycle greenhouse gas emissions of concrete production. a global average cubic meter of non-reinforced concrete is responsible for approximately 200 kg of co2-eq. [19]. this may be over 400 kg of co2-eq. for high-performance concrete (e.g., 50 mpa [19]). either way, the impact of desalination and water distribution for the concrete mix would be far less than 1% of the total impact. • greenhouse gas emissions for desalination are 1.5 kg co2-eq/m3 desalinated water. this is lower than the values reported for other geographical contexts. in qatar, for instance, 9.5 kg of co2-eq is estimated to be emitted per m3 of desalinated water [9]. there are two potential reasons: (i) only impacts associated with the electricity consumption of the desalination plant (3.5 kwh/m3 dsw) are considered in this study and, (ii) the lower greenhouse gas intensity of electricity used in israeli desalination plants. • the thresholds for reducing water and carbon footprints by using untreated seawater in the mix in place of municipal grid water are 13 and 22 km, respectively. if a batching plant is located less than 13 km away from the desalination plant (x < 13), transporting seawater by truck instead of desalinating it and pumping it through the national pipeline network would reduce both the water and carbon footprints of concrete. if the distance is between 13 and 22 km, desalinating and distributing water would have a lower water footprint but would generate higher greenhouse gas emissions. finally, if the batching plant is located more than 22 km away from the desalination plant, transporting seawater by truck would lead to higher impacts for both water and carbon. 7. conclusions the water and carbon footprint consequences of using seawater instead of municipal grid water for concrete mixing in israeli batching plants are estimated. a net benefit for both impacts would emerge only if the batching plant is located less than 13 km from a desalination plant. however, even in this case, the benefits appear to be negligible compared with the average life cycle impacts of concrete. negligible savings are observed since the marginal water supply in israel is desalinated seawater, which has a water footprint of zero and a low carbon footprint compared to other constituents of concrete, such as cement. nevertheless, if batching plants are located near the coast, given the large amount of concrete produced in israel each year, significant greenhouse gas savings could still be achieved: for every km less than the 22-km threshold, 20 g of co2-eq. could be spared for every m3 of concrete produced. if, theoretically, all israeli concrete was produced in plants that are less than 10 km away from desalination plants, over 3,000 tons of co2-eq. could be saved every year. moreover, by using untreated seawater in the mix, other environmental impacts (e.g., marine ecotoxicity, biodiversity) associated with the desalination process, which are not included in the present assessment, could be avoided. approximately 1.5 m3 of brine are produced for every m3 of desalinated water and concerns are growing over the rising toxic levels from its discharge back into the sea [8]. to confirm the results, a sensitivity analysis on the assumptions and data should be performed. for instance, results might differ if pipes were used instead of trucks to deliver the seawater or if seawater was to replace potable water used for washing the quarried aggregates as well. moreover, even though negligible savings (or larger impacts, when distances from the coast are above the threshold) are observed in israel, contrary conclusions may result in countries where the marginal supply of municipal grid water comes from natural freshwater resources. the impact of seawater on the service life of the machinery used for concrete production and of the concrete structure itself should be further investigated to avoid burden shifting. finally, a comparison of the life cycle costs of the two scenarios would provide valuable information to the concrete industry and the government to evaluate the practicality and economic viability of shifting the concrete industry towards the use of seawater. in a future analysis the challenges of a design practice not yet widespread and the market cost of gfrp should be included. references [1] m. m. mekonnen, a. y. hoekstra. four billion people facing severe water scarcity. science advances 2(2), 2016. https://doi.org/10.1126/sciadv.1500323. [2] y. satoh, t. kahil, e. byers, et al. multi-model and multi-scenario assessments of asian water futures: the water futures and solutions (wfas) initiative. earth’s future 5(7):823-52, 2017. https://doi.org/10.1002/2016ef000503. 30 https://doi.org/10.1126/sciadv.1500323 https://doi.org/10.1002/2016ef000503 vol. 33/2022 concrete made of desalinated water [3] d. rodríguez-robles, p. van den heede, n. de belie. life cycle assessment applied to recycled aggregate concrete. new trends in eco-efficient and recycled concrete, p. 207-56. https: //doi.org/10.1016/b978-0-08-102480-5.00009-9. [4] s. a. miller, a. horvath, p. j. m. monteiro. impacts of booming concrete production on water resources worldwide. nature sustainability 1(1):69-76, 2018. https://doi.org/10.1038/s41893-017-0009-5. [5] k. r. o’brien, j. ménaché, l. m. o’moore. impact of fly ash content and fly ash transportation distance on embodied greenhouse gas emissions and water consumption in concrete. the international journal of life cycle assessment 14(7):621-9, 2009. https://doi.org/10.1007/s11367-009-0105-5 . [6] v. arosio, a. arrigoni, g. dotelli. reducing water footprint of building sector: concrete with seawater and marine aggregates. iop conference series: earth and environmental science 323(1), 2019. https://doi.org/10.1088/1755-1315/323/1/012127 . [7] a. m. neville. properties of concrete fourth ed. (harlow, uk: longman group limited), 1995. [8] e. jones e, m. qadir, m. t. h. van vliet, et al. the state of desalination and brine production: a global outlook science of the total environment, 657:1343-56, 2019. [9] e. jones, m. qadir, m. t. h. van vliet, et al. the state of desalination and brine production: a global outlook. science of the total environment 657:1343-56, 2019. https: //doi.org/10.1016/j.scitotenv.2018.12.076. [10] m. mannan, m. alhaj, a. n. mabrouk, et al. examining the life-cycle environmental impacts of desalination: a case study in the state of qatar. desalination 452:238-46, 2019. https://doi.org/10.1016/j.desal.2018.11.017. [11] e. redaelli, a. arrigoni, m. carsana, et al. culvert prototype made with seawater concrete: materials characterization, monitoring, and environmental impact. advances in civil engineering materials 8(2), 2019. https://doi.org/10.1520/acem20180114. [12] f. lollini, m. carsana, m. gastaldi, et al. seawater and stainless steel bars for sustainable reinforced concrete structures. iop conference series: earth and environmental science 296(1), 2019. https://doi.org/10.1088/1755-1315/296/1/012017. [13] a. katz. personal communication, 2019. [14] t. ekvall, b. p. weidema. system boundaries and input data in consequential life cycle inventory analysis. the international journal of life cycle assessment 9(3):161-71, 2004. https://doi.org/10.1007/bf02994190. [15] y. parag. personal communication, 2019. [16] t. opher, e. friedler. comparative lca of decentralized wastewater treatment alternatives for non-potable urban reuse. journal of environmental management 182:464-76, 2016. https://doi.org/10.1016/j.jenvman.2016.07.080. [17] b. p. weidema, n. frees, a.-m. nielsen. marginal production technologies for life cycle inventories. the international journal of life cycle assessment 4(1):4856, 1999. https://doi.org/10.1007/bf02979395. [18] t. opher. a multi-objective lca-based model for sustainability assessment of urban water reuse alternatives at various centralization scales, department of civil and environmental engineering, technion israel institute of technology), 2016. [19] g. wernet, c. bauer, b. steubing, et al. the ecoinvent database version 3 (part i): overview and methodology. the international journal of life cycle assessment 21(9):1218-30, 2016. https://doi.org/10.1007/s11367-016-1087-8. [20] j. netz, j. sudin. water footprint of concrete, department of sustainable development, environmental science and engineering, kth royal institute of technology), 2015. [21] a.-m. boulay, j. bare, l. benini, et al. the wulca consensus characterization model for water scarcity footprints: assessing impacts of water consumption based on available water remaining (aware). the international journal of life cycle assessment 23(2):368-78, 2017. https://doi.org/10.1007/s11367-017-1333-8. 31 https://doi.org/10.1016/b978-0-08-102480-5.00009-9 https://doi.org/10.1038/s41893-017-0009-5 https://doi.org/10.1007/s11367-009-0105-5 https://doi.org/10.1088/1755-1315/323/1/012127 https://doi.org/10.1016/j.scitotenv.2018.12.076 https://doi.org/10.1016/j.desal.2018.11.017 https://doi.org/10.1520/acem20180114 https://doi.org/10.1088/1755-1315/296/1/012017 https://doi.org/10.1007/bf02994190 https://doi.org/10.1016/j.jenvman.2016.07.080 https://doi.org/10.1007/bf02979395 https://doi.org/10.1007/s11367-016-1087-8 https://doi.org/10.1007/s11367-017-1333-8 https://doi.org/10.14311/app.2022.33.0271 acta polytechnica ctu proceedings 33:271–276, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague study for re-asr behaviour of recycled concrete using asr generated concrete and considering countermeasure technology takeshi iyodaa, ∗, nobuhiro matsudab a shibaura institute of technology, 3-7-5 toyosu koto-ku, tokyo, japan b tokyo techno company, 3343 onoji-machi machida city, tokyo, japan ∗ corresponding author: iyoda@shibaura-it.ac.jp abstract. it is predicted that the concrete volume to be discarded will increase due to the renewal of the concrete structures and the increase in the returned concrete from construction site. the use of this concrete block as recycled aggregate is very important in terms of sustainability. however, recycled aggregate concrete is difficult to use due to large problems such as drying shrinkage and freeze-thaw action. on the other hand, the aggregate used for the raw concrete is often unknown, suggesting the danger of asr. there is no previous research about what kind of danger is caused when an aggregate having an asr risk is used as a recycled aggregate. therefore, in this study, recycled aggregate was made from raw concrete where asr occurred, after that recycled concrete was manufactured using this aggregate. after this, asr tests were conducted again for the recycled concrete. in addition, we examined whether recycled concrete using aggregates with properties improved by using the previously reported carbonation technique could be an asr countermeasure technology. as a result, it was found that the recycled concrete using the asr-generated raw concrete as aggregate can suppress asr by using recurring asr aggregate treated by carbonation technology. keywords: carbonation technology, re-asr behavior, recycled concrete. 1. introduction it is predicted that the concrete volume to be discarded will increase due to the renewal of the concrete structures and the increase in the returned concrete from construction site. it is important to use the recycled aggregate from this concrete to achieve sustainability society. further, in order to spread recycled aggregate concrete, it is desired to use lowquality recycled aggregate that also generates few byproducts such as fine powder and can be manufactured reducing energy and cost. however, recycled aggregate concrete has significant problems such as low strength, drying shrinkage and freeze-thaw. furthermore, the recycled aggregate reactiveness is often unknown for the danger of alkali-silica reaction (asr). it should be noted that in asr, alkali supply from attached mortar on recycled aggregate is also assumed. there has been little discussion about the dangers that would occur if concrete with reactive aggregates were used as recycled aggregate. in order to promote the use of low-quality recycled aggregate in the future, it is necessary to get more knowledge about asr. on the other hand, in past works, the authors have proposed a modification technique by carbonation for the purpose of improve low-quality recycled aggregate [1–4]. this technology focuses on the carbonation mechanism of concrete, it applies high-concentration carbon dioxide on the recycled aggregate to carbonate the mortar in the recycled aggregate, thereby modifying the recycled aggregate itself. it has been confirmed that recycled aggregate modified by this technique meliorate concrete strength, length change by drying shrinkage and freeze-thaw resistance. however, the impact on asr has not been studied. carbonation was expected to have the effect of lowering the alkali concentration of the mortar, and was thought to be helpful in suppressing asr. therefore, in this study, the raw concrete was prepared by using the reactive coarse aggregate in order to develop asr. the recycled coarse aggregate was manufactured from this raw concrete with confirmed asr, and the recycled concrete was prepared. at that time, recycled coarse aggregates low (l) and high (h) quality were produced. the recidivism of asr (re-asr) of this recycled concrete were experimentally investigated. in addition, the effect of asr suppression using modified recycled aggregate by carbonation was also investigated. 2. outline of the experiments 2.1. preparation of raw concrete and asr characteristics table 1 shows the types and physical properties of the aggregates. regarding the rock types of each coarse aggregate, og is crushed hard sandstone, kg is andesite, and sg is gravel. the asr characteristics of the aggregate used were confirmed by the chemical method (jis a 1145), the mortar bar method (jis a 1146) and the accelerated mortar bar method (astm 271 https://doi.org/10.14311/app.2022.33.0271 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en takeshi iyoda, nobuhiro matsuda acta polytechnica ctu proceedings kinds dry density water absorption ration f.m. sc rc classify of asr! g/cm3 " [%] [mol/l] [mol/l] os fine agg. 2.55 2.18 2.82 28 84 harmless og coarse agg. 2.63 0.89 6.62 40 78 harmless kg 2.72 1.72 6.41 388 94 not harmless sg 2.60 1.18 7.41 73 107 harmless table 1. the types and physical properties of the aggregate. figure 1. mortar bar test on jis. c 1567). the test results of the chemical method are shown in table 1. the os (fine aggregate), og and sg were judged as "harmless" and kg was judged as "not harmless". figure 1 shows the results of the mortar bar test using jis. all the coarse aggregates were judged to be "harmless" with less than 0.1% of expansion ratio at the age of 26 weeks. however, for kg, the expansion increased after 13 weeks of age, and it is estimated that the expansion further increased after 26 weeks. figure 2 shows the results of the accelerated mortar bar test by astm. the kg and sg were judged "harmful", and og was "harmless". it can be seen that the expansion amount of kg greatly exceeds 0.2% at 14 days of age, and the reactivity is high. the sg gradually expanded from around 5 days of age and exceeded 0.2% at 14 days of age, indicating that it is an aggregate with low to moderate reactivity. based on the above results, kg and sg were determined to be reactive aggregates, and og was determined to be harmless aggregates. using these aggregates, raw concrete was produced with the planned composition shown in the table 2. the cement used was ordinary portland cement with a na2o equivalent of 0.57%. in order to promote asr, naoh was added during mixing so the amount of alkali in concrete increased to 7.0 kg/m3. table 2 also shows the fresh properties of the raw concrete and the compressive strength when cured in tap water for 28 days. regarding the fresh properties and compressive strength of the raw concrete, there figure 2. accelerated mortar bar test on astm. was no significant difference between the slumps, although it varied slightly. however, the compressive strength was slightly lower due to on the effect of adding naoh. next, a rilem aar-3 (38◦c concrete prism test) test was performed to evaluate on the asr characteristics of the raw concrete. figure 3 shows the results of the rilem aar-3 test. the criteria for rilem aar-3 is described as those samples with a swelling above 0.04% at the age of 52 weeks are judged to be harmful. the kg aggregate concrete (kc) increases the swelling from around 10 weeks of age and 0.04% at the age of 52 weeks. cracks were confirmed at 14 weeks. the swelling of sc gradually increased around 10 weeks, and slightly exceeded 0.04% at 52 weeks. on the other hand, oc was determined to be "harmless." these results show the same tendency as the asr characteristics of the raw aggregate on the accelerated mortar bar test as astm. in order to prepare recycled coarse aggregate, cylindrical specimens of ! 100 × 200 mm were prepared. the raw concrete specimens were stored in an outdoor environment for one to three weeks to further promote asr, and then cured in water at 40◦c. to check the amount of expansion of the raw concrete specimen, a 100 × 100 × 400 mm prismatic sample was prepared. figure 4 shows the amount of expansion of the monitored specimens. the swelling behaviour of these samples were slower when compared with rilem aar-3 test. this is because the monitoring specimens were cured under the same conditions as 272 vol. 33/2022 re-asr behaviour of recycled concrete kinds of unit weight fresh concrete properties compressive coarse agg. w/c c w s g slump air temp. strength [%] [kg/m3] [cm] [%] [◦c] [n/mm2] oc og 981 11.0 4.7 25.0 29.4 kc kg 50 350 165 823 1021 6.5 4.3 24.0 28.7 sc sg 962 16.0 4.1 23.0 27.9 table 2. mix proportion, fresh properties and compressive strength of concrete at 28 days. figure 3. asr expansion test on rilem. the concrete specimens. however, the amount of expansion of kc increased rapidly from the age of 13 weeks, showing a similar tendency to the raw aggregate and raw concrete. at 20 weeks of age, cracks were confirmed in the test specimen. on the other hand, in sc, although the amount of expansion was slightly small, there was no significant difference from the expansion behaviour in the rilem aar-3 test. recycling of the original concrete specimens was performed at about 20 weeks for kc and 26 weeks for oc and sc at the age of the monitoring specimens. 2.2. production of recycled coarse aggregate recycled coarse aggregates l and h and modified recycled aggregate lc were produced from oc, kc and sc raw concrete specimens. at this point, kc has confirmed expansion and cracking due to asr. recycled coarse aggregate l was produced only by crushing treatment (jaw crusher / impact crusher). recycled coarse aggregate h was produced by further grinding (ball milling) the recycled coarse aggregate l. in the modified recycled aggregate, the coarse aggregate l was carbonated. the forced carbonation condition was 1 week long under the of temperature of 20◦c and relative humidity of 60% in the accelerated carbonation chamber, as in the previous study [1, 3]. all the produced recycled coarse aggregates were washed with water to remove fine particles. figure 5 shows the density and water absorption of the recycled coarse aggregate. in all types of modified recycled aggregate, the density and water absorption figure 4. expansion test on monitoring specimens. figure 5. density and water absorption ratio on recycled aggregates. properties had improved. there are also improvements for the quality of recycled coarse aggregate m. although the alkali content was separately measured, no significant difference was found in the total alkali content of the recycled coarse aggregates l and h in this study. 3. asr characteristics of recycled aggregate concrete 3.1. outline of the experiment concrete was prepared using the recycled coarse aggregate shown in table 3 and the asr characteristics were confirmed. the materials used the same as preparing recycled aggregate concrete. the composition of recycled aggregate concrete was w/c = 0.5, 273 takeshi iyoda, nobuhiro matsuda acta polytechnica ctu proceedings kinds of concrete coarse agg. compressive strength coarse agg. non-naoh adding naoh ol oc og 37.1 24.2 olc 38.5 27.5 oh 37.1 38.5 kl kc kg 39.3 25.5 klc 40.3 30.7 kh 40.1 27.3 sl sc sg 35.8 24.5 slc 38.5 26.7 sh 39.9 27.8 table 3. results of compressive strength on different concrete. s/a = 0.46, and unit water volume of 165 kg/m3. for the purpose of promoting asr, two types were prepared, one with naoh added and one without naoh, so that the total alkali content in recycled aggregate concrete including alkali from cement was 7 kg/m3. in addition, the calculation of the total alkali content of the sample to which naoh was added did not include the alkali content from the aggregate. a 100 × 100 × 400 mm prism was prepared, and after demolding, it was cured for 1 week in an environment at a temperature of 20◦c and humidity of 60%. table 3 shows also the compressive strength on recycled concrete. the compressive strength tends to be larger for h class aggregate. it can be said that the amount of attached mortar affects the strength development. also, the modified recycled aggregate has increased compressive strength compared to l, as in previous studies [2, 4]. when naoh was added, the compressive strength was reduced. 3.2. experimental results figure 6 shows the expansion of recycled aggregate concrete. comparing the types of raw aggregates, those using the non-reactive raw aggregate og showed almost no expansion regardless of whether naoh was added or not. it is clear that there is no danger of asr occurring again when recycled coarse aggregate made from non-reactive raw aggregate is used. in the case of using highly reactive kg, the amount of expansion was significantly increased with the addition of naoh. silica remains at the time of production of recycled coarse aggregate, and it is considered that asr occurred when recycled aggregate concrete was used. in addition, those without addition of naoh showed almost no swelling. this is probably because the total alkali content of recycled aggregate concrete using kg and not containing naoh was 2.9 kg/m3, indicating that the alkali content did not reach asr. on the other hand, when sg was used, the aggregate had low to moderate reactivity as the asr characteristics of the original aggregate, but almost no expansion was observed as in og. comparing l and h, h increased the rate of expansion when naoh was added and using reactive aggregate. this is considered to be due to the fact that the amount of attached mortar with h was small, so that the elution of silica and the penetration of alkali were easily performed. on the other hand, in the case of using kg with high reactivity without addition of naoh, l slightly expanded. it is possible that the alkali in the attached mortar affects asr. in addition, the amount of expansion is greatly suppressed by using the modified recycled aggregate. in particular, those using highly reactive kg as the raw aggregate were very remarkable, showing a high suppression effect on asr. figure 7 shows an image of this asr suppression mechanism. in addition to the presence of attached mortar, the modified recycled aggregate become denser by carbonation, which suppresses the penetration of alkali and further dissolution of silica in the raw aggregate. it is thought that asr is reduced by suppressing it. however, in this experiment, the expansion of the original concrete due to asr was considered to have calmed down to a certain extent, and the scope of the experiment was limited. in the future, it is necessary to continue studying the raw concrete before asr progresses, and also to confirm the suppression mechanism by the modified recycled aggregate. 4. summary the following findings were obtained in this study. 1. the asr characteristics by the accelerated mortar bar method in raw aggregate and rilem aar-3 in raw concrete showed the same tendency. 2. it was found that there was no danger of asr occurring again when recycled coarse aggregate made from non-reactive raw aggregate was used. in the case of using highly reactive kg, the amount of expansion significantly increased with the addition of naoh. this suggests that even when the asr of the raw aggregate has progressed to some extent, the silica remains when the recycled coarse 274 vol. 33/2022 re-asr behaviour of recycled concrete figure 6. results of re-asr on recycled concrete. figure 7. the mechanism of prevented asr on using carbonated recycled aggregate. 275 takeshi iyoda, nobuhiro matsuda acta polytechnica ctu proceedings aggregate is made, and when the recycled aggregate concrete is used, the asr occurs again. 3. when l and h quality recycled aggregate were compared, the expansion was higher for h with the addition of naoh and using reactive aggregate. 4. the swelling behaviour was greatly suppressed by using the recycled aggregate modified by carbonation, and for the one using highly reactive kg as the original aggregate the decrease was very remarkable, showing a high inhibitory effect on asr. acknowledgements this experiments were carried out by mr. abdulkareem abdulkadeer abdullahi and the all members of our laboratory. references [1] n. matsuda, t. kameyama, m. matsuda, et al. study on the production method of low-energy recycled aggregate by forced adsorption of co2 gas. proceedings of the japan concrete institute annual convention 36:1732-1737, 2014. [2] n. matsuda, s. suzuki, t. iyoda. proceedings of the japan concrete institute annual convention 38:1791-1796, 2016. [3] t. iyoda, n. matsuda. ii international conference on concrete sustainability, pp. 132-141, 2016. [4] a. a. abdulkadeer, n. matsuda, t. iyoda. ialcce2018 the sixth international symposium on life-cycle civil engineering, pp. 575-582, 2018. 276 https://doi.org/10.14311/app.2022.33.0337 acta polytechnica ctu proceedings 33:337–343, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague influence on permeability and pore structure of polyolefin fiber reinforced concrete containing slag wei-ting lina, ∗, an chenga, kinga korniejenkob, michał łachb a department of civil engineering, national ilan university, taiwan b institute of materials engineering, faculty of mechanical engineering, cracow university of technology, poland ∗ corresponding author: wtlin@niu.edu.tw abstract. the purpose of this study is to assess the mechanical and microscopic properties of concrete containing different dosages of polyolefin fibers and slag through tests of compressive strength, resistivity, water absorption, mercury intrusion porosimetry and scanning electron microscopy. test results indicate that the specimens containing slag have higher compressive strength, lower absorption, lower resistivity and denser porestructures than the control and specimen made with fibers. the specimens containing slag and polyolefin fiber demonstrated better performances in fiber reinforced concrete. scanning electron microscopy illustrates that the polyolefin fiber acts to arrest the propagation of internal cracks. still, there are cracks and weaknesses between fiber and paste that cause harmful ions penetrated easier. keywords: critical pore size, mercury intrusion porosimetry, slag. 1. introduction concrete is a porous and cracked brittle material. it has low tensile strength, poor toughness, and small ultimate tensile strain, so it is easy to crack. the existence of cracks in the concrete would reduce the structural capacity and increase the deflection. at the same time, external harmful ions would penetrate into the concrete through the cracks, eroding the concrete or corroding the internal rebars, which would seriously affect the durability of the concrete structures. especially in modern concrete engineering, buildings are developing in the direction of large volume, large area, and complex and diverse shapes, and then concrete is developing in the direction of high strength and superior fluidity. therefore, it was required that the strength and slump of the concrete continue to enhance, and the amount of cement continues to increase. the negative impacts were that the hydration heat, shrinkage, shrinkage deformation stress and the number of cracks of concrete would be increased [1]. in order to solve the cracking problem, researchers have adopted the method of adding fibers to concrete to improve its cracking, thereby increasing the durability of the concrete. on the other hand, by reducing the water-tobinder ratio of concrete or adding pozzolanic materials, such as fly ash, ground-granulated blast-furnace slag (slag), silica fume, etc., it can help improve the pore structures in concrete [2–4]. shannag was suggested to use a certain type of fly ash mixed with silica fume and replace partial cement to improve the compressive strength, splitting strength and elastic modulus of concrete. relevant research has also pointed out that good concrete materials should have several excellent material properties, including workability, strength, corrosion resistance and high resistance to ion penetration [5–7], and the addition of pozzolaic materials can meet the above requirements, providing better macroscopic and microscopic properties of concrete. the most effective method for evaluating the macroscopic properties of concrete was the compressive strength test. mercury intrusion porosimetry (mip) and scanning electron microscopic (sem) observation were needed to understand the microscopic part of the concrete. it can be assessed the strength properties and transmission characteristics of durability on concrete structures. studies in recent years have found that to understand the pore-structures and transmission characteristics of concrete, porosity and pore size distribution are important research parameters. porosity and pore size distribution are the most direct parameters for the permeability and pore structures of cement pastes. permeability is directly related to the continuity of fluid passage (pore diameters of 120 or 160 nm) [8]. porosity is a measure of the ratio of pores to the total volume of the pastes. if the specimens had a high porosity and the pores were connected, the cement-based materials also had higher permeability. on the contrary, if the pores were not connected, the permeability of the materials was tended to decrease. it can be seen that the connectivity of the pores had a considerable relationship between permeability and durability [9, 10]. as far as the pore structures of concrete were concerned, even if there 337 https://doi.org/10.14311/app.2022.33.0337 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en w. lin, a. cheng, k. korniejenko, m. łach acta polytechnica ctu proceedings chemical compositions (%) materials cao sio2 al2o3 fe2o3 mgo so3 k2o l.o.i. cement 63.9 20.7 5.4 3.2 2.0 − 0.7 0.05 slag 41.6 33.5 14.1 0.4 6.9 0.7 2.2 0.78 table 1. chemical compositions of slag and cement. mix no. water cement slag fine aggregates coarse aggregates fibers a 217 395 0 908 780 0 ap1 395 0 902 774 3.6 ap2 395 0 897 769 7.3 ap3 395 0 892 764 10.9 ap4 395 0 887 759 14.5 ap5 395 0 882 754 18.2 ag4 237 158 908 780 0 ag4p1 237 158 902 774 3.6 ag4p2 237 158 897 769 7.3 ag4p3 237 158 892 764 10.9 ag4p4 237 158 887 759 14.5 ag4p5 237 158 882 754 18.2 table 2. mix design of concrete (kg/m3). test target specimen dimensions(mm) referenced standard mechanical properties compressive strength test !100 × 200 astm c39 permeability absorption !100 × 200 astm c642 micro-structure observations mip 10 × 10 × 3 astm d4404 sem observation 10 × 10 × 3 astm c1723 table 3. mix design of concrete (kg/m3). were more other characteristics that would affect the behavior and other engineering characteristics. the porosity and pore size distribution were important pore-structure parameters, which would directly affect the strength, durability and permeability of the concrete. this study is aimed to evaluate the influence on permeability and pore-structures of polyolefin fiber reinforced concrete containing slag using compressive strength, absorption, mip and sem observations. 2. experiments 2.1. materials in this study, it used type i portland cement with the specific gravity of 3.15 and the fineness of 3310 cm2/g. the fine aggregate was natural river sand with ssd specific gravity of 2.56, the absorption was 2.25% and the fineness modulus was 2.82. the coarse aggregates had an ssd specific gravity of 2.65 and absorption of 1.44%. the ground-granulated blast-furnace slag (g) with a white powder was produced by chc resources corporation in taiwan and used as constituents of cement replacement. slag had the specific gravity of 2.88 and fineness of 4000 cm2/g. the pozzolanic activity index of slag at the age of 28 days was 127%. the chemical compositions of slag and cement were summarized in table 1. inclusion of polyolefin fibers in composites can enhance the tensile strength and volume stability of cement-based materials containing slag. the length and aspect ratio of the fiber is 25 mm and 200 (d = 0.125 mm), respectively. the specific gravity, tensile strength and young’s modulus of fiber is 0.90, 275 mpa and 2647 mpa, respectively. 2.2. mix design and test methods the water/cement ratio (w/c) of the concrete specimens was maintained at a constant 0.55 in accordance with the aci 211.1 specification. table 3 lists the mix design of mixtures for the concrete specimens. the specimens were numbered using three letters and numbers to indicate the slag replacement of cement and dosage of the polyolefin fibers. a denotes ordinary portland concrete; p1, p2, p3, p4 and p5 refers to specimens containing 0.4 vol.%, 0.8 vol.%, 1.2 vol.%, 1.6 vol.% and 2.0 vol.% fibers, respectively. g4 refers to specimens containing 40% slag. table ?? presents the tests performed, the dimensions of the specimens and the standards used in this study. 338 vol. 33/2022 analysis of concrete and cement epd figure 1. compressive strength development curves of polyolefin fiber concrete without slag. 3. results and discussion 3.1. compressive strength figures 1 and 2 presents the compressive strength of polyolefin fiber concrete made with and without slag at the age of 7, 28, 56 and 91 days, respectively. the compressive strength increased significantly with the increase in the amount of fibers at 7 days of test age. it indicated that the addition of fibers could reduce the autogenous shrinkage or fine cracks of the specimens and the fibers were played a very important role in the early stage of crack resistance. the strength development and growth rate gradually flattens at the age of 28 days and the effect of age on compressive strength gradually decreases, and the amount of fibers had no significant effect on compressive strength (increasing trend was the same for 56 to 91 days). comparison of different fiber additions, the specimens containing 2.0% fibers had lower compressive strength than that of the specimens containing 1.6% fibers at different ages. it might be due to the addition of a higher fiber addition amount and it might affect the workability of the specimens. the worse workability was caused the fiber to be difficult to be uniformly dispersed in the specimens, which resulted in a lower compressive strength [11]. test results of the specimens with and without slag found that the compressive strength of the specimens increased with the increase of curing age, and the specimens containing slag had higher compressive strength than that of specimens without slag. the compressive strength was tended to increase with the increase of the inclusion of fibers. comparing test specimens at different ages, the specimens containing 2.0% had a slightly higher compressive strength than the specimens containing 1.6% fibers. it indicated that the inclusion of slag in concrete could increase the compressive strength of fiber concrete. the inclusion of slag in the specimens was reflected better workability and enabled the fibers to be uniformly dispersed during the mixing process, resulting in a figure 2. compressive strength development curves of polyolefin fiber concrete with slag. better compressive strength. at the same curing age, the compressive strength of the polyolefin fiber reinforced concrete with 40% slag (ag4p5 specimens) could be 12% higher than the ap5 specimens. fibers also can improve the bonding strength of the fiber interface, and then increase the compressive strength of the concrete. 3.2. absorption absorption is an important indicator for judging the permeability and quality of concrete. lowpermeability concrete can resist the penetration of water, and it also made other harmful ions difficult to penetrate into the concrete. a comparison of the influence on the absorption of the addition of 40% slag on polyolefin fiber reinforced concrete at various age is shown in figures 3 and 4, respectively. it indicated that the absorption was decreased with the increase of curing age, resulting in denser pore-structures. however, the absorption of each group of specimens without slag was remained at about 6%, which did not vary with the inclusion of fibers in concrete. this might be caused by the addition of polyolefin fibers in the specimens and the fibers could reduce the connectivity of the pores. in addition, the weak interface of the fibers had a greater impact on the absorption, resulting in varied values of absorption [12]. the results found that the specimens mixed with 40% slag and 0.4% polyolefin fibers were reflected in lower absorption of various ages. as the amount of fiber increased, the absorption was started to increase. it was obvious that even if slag was added to improve the adhesion between the pastes and the polyolefin fibers. as long as the fibers were slightly unevenly dispersed, it would affect the results of absorption. it was necessary to ensure that the blended polyolefin fibers were evenly distributed in the specimens and prevented the fiber from agglomerating to ensure that it had no agglomerate seriously, it makes poor mechanical properties and permeability of the 339 w. lin, a. cheng, k. korniejenko, m. łach acta polytechnica ctu proceedings figure 3. absorption development curves of polyolefin fiber concrete without slag. figure 4. absorption development curves of polyolefin fiber concrete with slag. concrete. compared with the control specimens, the combination of 0.4% polyolefin fibers and 40% slag reduced the absorption of the specimens. however, the maturity of the specimen had a limited effect on absorption. the polyolefin fiber-reinforced concrete containing 40% slag had lower absorption, which was 50% and 35% lower than the a and ap1 specimens. the fiber could reduce the connectivity of the pores, and the slag could strengthen the degree of adhesion between the fiber and the paste to achieve the double enhancement effect. figure 5 is illustrated the relationship between the absorption and the polyolefin fiber content of each mix of polyolefin fiber-reinforced concrete containing slag at the age of 91 days (testing results were averaged by three specimens for each mixture and the coefficient of variation was controlled under 5%). from the regression equation, the absorption of fiber-reinforced specimens mixed with 40% slag was changed with the increase of fiber addition. when the figure 5. relationship between absorption and fiber comtent (40% slag). figure 6. relationship between the pore size and cumulative mercury intrusion of polyolefin fiberreinforced concrete specimens (without slag). fiber amount reached 1.55%, there was the highest absorption; however, the fiber amount reached 0.45%, there was the lowest absorption. its tendency was similar to that of the specimens without slag. the results also indicated that the specimens containing 40% slag had a lower absorption than that of specimens without slag. 3.3. mip results figures 6 and 7 show the relationship between the pore size and cumulative mercury intrusion of the polyolefin fiber-reinforced concrete specimens with and without slag at the age of 91 days. the test results found that the cumulative mercury intrusion and total porosity of the specimens decreased with the increase of polyolefin fibers. the cumulative mercury intrusion of ap5 specimens was 0.0654 ml/g, which was lower about 43% than the control specimen (a specimen was 0.1147 ml/g). the lowest cumulative mercury intrusion was 0.0288 ml/g of ap2 specimens, although it was lower than that of other 340 vol. 33/2022 analysis of concrete and cement epd figure 7. relationship between the pore size and cumulative mercury intrusion of polyolefin fiberreinforced concrete specimens (with slag). groups. the ap3 specimen had a high cumulative mercury intrusion (0.1288 ml/g), which might also be caused by the weak interfacial transition zone between the pastes and the fibers described in the previous section, but overall it has appeared that the addition of polyolefin fibers was helped to reduce the internal porosity of specimens. from figure 6, the pore distribution curve of the specimen had a relatively gentle trend with the increase of fibers (especially when the pore size is less than 50 nm). it can be seen that the addition of fibers can also help reduce the pore size to a certain extent. it was also verified that the fiber can reduce the connectivity between the pores and achieve the effect of improving the permeability of the concrete [13–15]. figure 7 shows that the cumulative mercury intrusion and total porosity of the specimens had a tendency to decrease significantly with the increase of the slag. the cumulative mercury intrusion of ag4 specimens was 0.0534 ml/g, which was reduced by about 53% than that of control specimens. however, the cumulative mercury intrusion and total porosity of the specimens mixed with slag were tended to decrease with the increase of polyolefin fibers. it was still lower than that of various groups of polyolefin fiber reinforced concrete specimens without slag. it can be seen that the inclusion of slag was more effective in reducing pores and reducing pore size than the addition of polyolefin fibers. except for the ag4p1 specimens, the pore distribution curve of all specimens had a gentler trend with the increase of fibers (especially when the pore size was less than 50 nm). it was found that the addition of fibers can reduce the pore connectivity. however, the effect of the addition in fiber-reinforced specimens on increasing compactness and reducing porosity was significant. it has been verified that the addition of slag in concrete can cause secondary hydration reaction with cement from ca(oh)2 to produce denser c-s-h colloid, thereby reducing the amount of free ca(oh)2, and using cfigure 8. sem photo of the control specimens. s-h colloid to reduce the volume and size of pores, thereby achieving the effect of improving the permeability of concrete [16]. the specimens combined with slag and polyolefin fibers might reflect lower pore characteristics than the control group or the specimens containing polyolefin fibers singly. however, the enhancements of the specimens combined with slag and polyolefin fibers were also lower than those of the specimens containing slag singly. it can be seen that although fibers can reduce the connectivity between pores, the formation of the interfacial transition zone was prone to generate new connected pores [17]. the addition of slag did not completely improve this weak interfacial transition zone. compared with previous studies [4, 12, 17, 18], if this weak interfacial transition zone can be mixed with finer silica fume, it might be effectively improved and reduced the porosity and pore size significantly. 3.4. sem observations figures 8 to 9 are shown the sem observations of control specimens and the specimens containing 40% slag, respectively. figure 8 indicated that more pores can be observed on the surface of the control specimens. the unevenness of the surface on pore structures was reflected in looser pore structures, which in turn results in concrete with lower mechanical properties and higher absorption. it was obvious that the microscopic properties had a considerable effect on the properties of the concrete including strength and permeability. the inclusion of slag in concrete can effectively increase the compressive strength of the concrete and improve permeability. the main reason is that the secondary hydration reaction produced by slag can provide the denser microstructures and the results can be seen from figure 9. it found that the pores had a significant decrease and were filled by hydration reactants. it was also verified that the addition of slag had a positive effect on concrete. figure 10 shows a surface observation diagram of 341 w. lin, a. cheng, k. korniejenko, m. łach acta polytechnica ctu proceedings figure 9. sem photo of the specimens containing 40% slag. polyolefin fiber. it can be found that the fiber surface is not as smooth as the original appearance. it is found that the surface was covered with fine hydration reactants and it was also showed an irregular surface of the fiber. these hydrates can increase the bonding force between the fiber and the pastes, resulting in an increase in the mechanical properties and permeability of the concrete. figure 11 shows the sem photo of the specimen containing 1.2% polyolefin fiber reinforced concrete and 40% slag. it was found that the inclusion of slag in fiber-reinforced concrete had a great benefit for the microstructures. the specimens containing slag had fewer capillary pores and were filled with the hydration reactants between the pores and the pastes. the findings correspond to the test results described in the previous section. 4. conclusions in terms of the results of compressive strength and absorption, the effect of the inclusion of fibers in concrete on the absorption and strength was limited. with the increase in the amount of fiber, the absorption was maintained at about 6%. the inclusion of slag in concrete can increase the strength and reduce the absorption of the specimens. comparing the specimens mixed with fiber and slag, there was no obvious trend in absorption. it is because of the interface between the fiber and the pastes caused by the weak interfacial transition zone. the weaker interfacial transition zone resulted in higher absorption; however, the ag4p1 and ag6p2 specimens still had lower absorption. regression analysis was performed on the compressive strength and the amount of fiber of the specimen, and it was found that there was a good trend relationship between them. it indicated that the best compressive strength was obtained from the specimens containing 1.35% fiber without considering the slag addition; however, the specimens containing 1.70% fiber had the best strength at the group of 40% slag specimens. figure 10. sem photo of the polyolefin fiber. figure 11. sem photo of the specimens containing 1.2% fiber and 40% slag. the sem and mip results can confirm that the addition of slag provided a denser microstructure of concrete, which improves the mechanical properties and permeability. fibers can provide the effect of inhibiting crack extension, and the interfacial transition zone between the fiber and pastes was prone to cause water or harmful ions to penetrate into concrete. the fibers were liable to produce a hydration reaction with the hydrates. the inclusion of slag in concrete can also reduce the weak interfacial transition zone, thereby achieving the effect of enhancing concrete performance and compactness. acknowledgements this work has been financed by polish national agency for academic exchange under the international academic partnership programme within the framework of the grant: e-mobility and sustainable materials and technologies emmat. this work also has been supported by the ministry of science and technology (most-109-2221e-197-001-my2) in taiwan. references [1] c. cao, m. m. s. cheung, b. y. b. chan. modelling of interaction between corrosion-induced concrete cover crack and steel corrosion rate. corrosion science 342 vol. 33/2022 analysis of concrete and cement epd 69:97-109, 2013. https://doi.org/10.1016/j.corsci.2012.11.028. 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https://doi.org/10.1016/j.cemconcomp.2004.09.010 my title https://doi.org/10.14311/app.2022.36.0244 acta polytechnica ctu proceedings 36:244–252, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague influence of plastic tension chord deformation on the shear capacity of prestressed beam elements sebastian thoma∗, oliver fischer technical university of munich, chair of concrete and masonry structures, theresienstraße 90, n6, 80333 münchen, germany ∗ corresponding author: sebastian.thoma@tum.de abstract. the shear strength of prestressed beam elements is of great interest with regard to the evaluation of existing concrete bridges. accompanying extensive experimental test series, the adaptation and modification of existing models for the evaluation of the load-bearing capacity also represents an important step. against this background, considerations are made for the isolation of critical influencing factors on the effective concrete compressive strength and quantified on the basis of a numerical framework. subsequently, a sensitivity analysis is performed on this basis to determine the influence of individual parameters and their interaction. taking into account the non-linear relationships at the cross-section level and in the cracked compressive stress field in the web, the shear reinforcement ratio and the strain in the longitudinal reinforcement are of major importance. keywords: plastic tension chord, prestressed concrete, sensitivity analysis, shear strength, stress fields. 1. introduction for the assessment of existing bridge structures and with respect to special questions related to prestressed concrete bridges, theoretical and experimental efforts to evaluate the shear capacity have been increased in the recent past [1–3]. in particular, characteristic bearing mechanisms of prestressed continuous girders with a low amount of shear reinforcement are continuously under discussion. from own experimental investigations, which are briefly presented schematically in section 2, the question for the evaluation of possible influencing factors on the effective concrete compressive strength of the cracked compressive stress field is developed and described by means of a numerical model, cf. section 3. subsequently, the proportion of the variance of the individual input variables in the total variance of the response spectrum is evaluated using a global sensitivity analysis [4] in section 4. from this, dominant parameters and model-inherent interactions can be derived. section 5 discusses main results and some influences of the numerical model are evaluated. existing approaches to consider the phenomenon of compression softening together with the presented results may be viewed from a new angle. this procedure provides a valuable basis for possible model modifications based on it. 2. experimental evidence the aim of the experimental research is the systematic illustration of the influence of a gradually reduced degree of longitudinal reinforcement and the associated effects on internal equilibrium. for this purpose, nine prestressed beam elements are tested using the substructure technique. figure 1 illustrates the concept. this method has already been deployed at the institute to study various aspects in regard to the shear strength of prestressed beam elements [1, 5], while the basic idea for testing detached subsystems has already been used successfully before [6]. including digital image correlation and fibre-optic sensors along longitudinal reinforcement and tendon axis, the experimental results show that the force flow in prestressed beams is subject to a highly statically indeterminate interaction that adopts a static equilibrium depending on communicating damage processes. in this way, the shear capacity is not considered in an exposed manner, but is to be classified with conformity of the stresses and strains under all acting internal forces, whereby in particular the mixed reinforced tension chord of reinforcing steel and tendon controls the cross-sectional strain and dominance of individual load-bearing components. the analysis shows that even with yielding longitudinal reinforcement under high bending moments in field and support regions, the ultimate system capacity is determined by a shear failure as long as the prestressing steel can provide the corresponding increase in strain. the evaluation of the tension chord deformation and the evolution of the cracked stress field in the web prove the great importance of arching actions in the description of the bearing behaviour of prestressed beam elements. a comprehensive analysis and discussion of experimental results will be presented in other publications and are not the subject of this paper. the brief outline of the experiments carried out is intended solely as a background and forms the basis for the considerations set out below, which can assist in evaluating the results and further refining the model concepts. 244 https://doi.org/10.14311/app.2022.36.0244 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 36/2022 influence of plastic chord deformation on shear capacity m v 𝜀sl a) b) c) fce = f(𝜀1) figure 1. experimental setup and focus of presented studies: a) substructure technique for investigating the shear strength of prestressed beam elements, based on the shear field of a continuous beam between the point load in the span and the inner support. reinforcement and tendons are anchored in tension in the cut edges/load plates; b) deformed system and internal forces acting on it. the push cam profiling enables the transmission of shear forces in analogy to bridges in segmental construction. the fans of the cracked compressive stress field are indicated in the field and support areas, a deflection takes place at tendon axis. the effective web compressive strength, controlled by the reduction factor kc, is the focus of the presented investigations. relevant details related to this factor are explained in the following section 3; c) qualitative profile of bending moment and shear force in the tested beam element. the load is successively increased until a shear failure occurs. under increasing load and a considerably reduced degree of longitudinal reinforcement, plastic strains result in the longitudinal reinforcement, which require a corresponding increase in strain in the tendons in order to guarantee the equilibrium of the internal forces. 245 sebastian thoma, oliver fischer acta polytechnica ctu proceedings 3. compression softening approach 3.1. background taking into account the test results, which show that the system load-bearing behaviour of prestressed crosssections is still limited by the shear capacity despite plastic tension chord deformation, and based on assumptions of the plasticity theory for the prediction of the shear strength, the effective concrete compressive strength of the cracked stress field is of crucial importance. as stated in [7], the assumption of a shear force bearing capacity derived from shear reinforcement is always limited by the bearing capacity of the concrete compression struts, which in turn depends on the effective concrete compressive strength and thus on the general state of strain. as a result, in addition to the shear reinforcement ratio, the longitudinal reinforcement ratio and the longitudinal strain coupled by the load distribution and system slenderness may have a major influence on the shear force bearing capacity. the factor kc, see equations 4 and 5, thus decides on the extent to which plasticity-theoretical approaches can be used to describe the shear strength of reinforced concrete beams. for design purposes, the effective concrete compressive strength is usually determined using constant reduction factors, which is a simple, conservative estimate. in the course of one’s own considerations, and especially in the case of plastic chord deformation, it is reasonable to explicitly take into account the actual state of strain. concrete girders, and in particular prestressed systems, develop a varying strain over beam height. the development of arching actions and the deviation of the stress field at the tendon axis cannot be represented in panel tests, which have contributed significantly to the description of various compression softening approaches [8–10]. still, the relations within the membrane element can be transferred to the beam web if an equivalent measure of the longitudinal distortion can be formulated. following the convention in the fib model code 2010 [11] or respectively the modified compression field theory (mcft) [9, 12], the measure of effective longitudinal strain is defined at the mid-depth of the member. however, the member is not defined by the cross-sectional contour, but must be understood as a web (following a classical plasticity-theoretical crosssectional division) between compression and tension chord. the longitudinal strain at half the height of the inner lever arm of the chord forces may serve as a clearer formulation here. 3.2. numerical model for the weighted evaluation of dominant influencing factors on the reduction factor kc and consequently on the effective concrete compressive strength, various assumptions are linked to a numerical model, the solution of which is solved as an optimisation problem, using sequential least squares [13, 14]. by iterating the strain plane for a cross-section under an acting bending moment and normal force, the lever arm of the internal forces is given after integrating the concrete compressive stresses, cf. figure 2. at the level of zm/2, the average longitudinal strain state is obtained, which is further supplemented by components caused by shear, cf. equation 1, adopted from [7]. 𝜎c δ𝜎p 𝜎s 𝜀c zm/2 zm/2 fc ft 𝜀x,mn800 250 figure 2. investigated cross section [mm], strain, stress, inner forces and decisive strain εx,mn due to bending and axial forces 𝜀z 𝜀x 𝜀2 𝜀1 𝜀 𝛾/2 𝜃 kc 𝜀1 a) b) figure 3. core components of the optimisation problem: a) mohr’s circle of strains in pure shear b) reduction factor kc, introduced equation according to [10, 15] since the proportion of longitudinal strain due to shear εx,v is already formulated as a function of the compressive strut angle θ, equation 1 is already part of the constrained nonlinear optimisation. based on equilibrium conditions and an explicit consideration 246 vol. 36/2022 influence of plastic chord deformation on shear capacity of the strains in longitudinal and shear reinforcement, which limit the range of possible strut angles (for a more detailed discussion, please refer to [7, 15]), the correlated values of the compression strut inclination and reduction factor kc, respectively the effective compressive strength fce, can be determined, cf. equations 4 and 5. assuming θ can evolve freely, taking into account its boundary conditions (derived from the yield strength and tensile strength of the reinforcement), the angle can be determined according to equation 2. the iterative routine is obvious in linking equation 2 and figure 3. after convergence the further steps are straightforward, and the principal strain ε1 as well as the reduction factor kc are determined, cf. equations 3 and 4. εx,v = v /2 · cot θ 2 (zs/zm · esas + zp/zm · epap) (1) tan θ = √ ρswfy fce − ρswfy (2) ε1 = εx,mnv + (εx,mnv − εc2) · cot2 θ (3) kc = 1 1.2 + 55 · ε1 ≤ 0.65 (4) fce = kc · ( 30 fc )1/3 · fc (5) the model uses additional, generally accepted simplifications, which are essentially due to the basic outlines of plasticity theory and will not be the focus of further attention here. related to the proposed model structure, some questions with respect to model sensitivity arise: • the principal strain εc2 is assumed to be constant at -0.002. is this, essentially empirically based, assumption [9] valid or do implicitly considered influencing factors falsify a generally valid statement? • what are the consequences of an increased strain level in the longitudinal reinforcement for established models for predicting the shear strength in prestressed beam elements? is εx a sensitive parameter for the response surface y? • are there significant interaction effects between different input parameters of the model? 4. global sensitivity analysis to address the above questions, the numerical model from section 3.2 is subjected to a sensitivity analysis. a global sensitivity analysis (gsa) distils the influence of possible uncertainty of selected n input parameters on the result of a model. possible interactions of the input values (so-called second order effects) can also be determined by variance-based methods. in the following sobol’s method [16, 17] enables a quantified evaluation of these issues. essential information on the characteristics of the gsa approaches used in this research is based on the considerations in [4], unless otherwise referenced. the method does not require an analytical model, but forms its correlation ratio on the basis of a problem spectrum to be defined initially, which contains all input parameters (which may be subject to the gsa), their distribution function and possible bounds. this data set is mapped into n dimensions by means of low-discrepancy sampling and then passed individually to the numerical model. critical aspects and details of the implementation are described below. 4.1. sobol’s sequence aiming for minimized discrepancy, generated sequences of parameter values should avoid holes and clusters in the hypercube. this task is adequately accomplished by sobol’s sequences [16, 17]. in addition, sobol’s sequence ensures order independence for different input parameter axes. without going into further detail, this is enabled by its direction numbers which are derived from distinct primitive polynomials for every required parameter vector [18]. sobol’ sequences are a quadrature rule and become unstable if samples of a size that is not a power of 2 are used, or if the first point is omitted, or if a sequence becomes sparse [19]. this method thus offers a tried and tested approach for evenly covering possible parameter constellations, which essentially eliminates the danger of unfavourable density of a parameter constellation. figure 4 shows an example of uniformly distributed random coordinates and the spread using a sobol’ sequence. the difference in the uniform coverage of the area is obvious. 0 1 0 1 0 1 0 1 0 1 figure 4. random samples and sobol’ sequence in two dimensions (27 = 128 samples) the model study accounts for seven parameters, which are listed in table 1 together with their assumed 247 sebastian thoma, oliver fischer acta polytechnica ctu proceedings uniform distribution. uniform distributions are assumed for the parameter space because no specific configuration with known uncertainty is to be examined, but an overall impression and overview of the model sensitivity is to be given. the limit values of the distributions are based on the test programme referred to at the beginning or assumed as realistic limit values when considering existing reinforced concrete bridges. all parameters are part of the resistance side and are included in the iteration of the initial strain plane and/or subsequent iteration to determine the stress and strain state in the web. in addition to the range of the longitudinal reinforcement ratio, the lower limit of which leads to plastic chord deformations in the crosssection analysis, the shear reinforcement ratio is also chosen relatively low with regard to the conditions in existing prestressed concrete bridges. taking into account the applied variation of the concrete compressive strength, a bandwidth of about 0.7 to 4.0 times the minimum shear reinforcement ratio ρw,min results for the general case according to ρw,min = 0.16ḟctm/fyk [20]. only λ represents a component of the load side to a certain extent, since the ratio of acting moment, which is increased independently of the gsa for all combinations of possible input values, and shear force, which also has a share in εx (see equation 1), is controlled via a variation of the shear slenderness. in addition, the variation of the tendon slope influences the effective shear force (after subtracting the vertical component vp = pmt · sin αp) in the web and ultimately the proportion εx,v. parameter uniform distr. range ρsl [%] [0.8, 2.0] ρsw [%] [0.09, 0.36] λ [-] [2.5 5.0] εc2 [-] [-0.002, -0.001] fc [mpa] [30.0, 50.0] σp [mpa] [550.0, 600.0] αp [deg] [1.0, 7.0] table 1. input parameters for sensitivity analysis. variance-based methods are powerful in quantifying the relative importance of input factors or groups. the main drawback of variance-based methods is the cost of the analysis, which, in the case of computationally intensive models, can become prohibitive even when using the approach described above. with saltellis extension of the sobol’s sequence [17], which is used within this scope, the resulting matrix has n (̇2n + 2) rows, where n is the number of parameters. for a full set of sobol’ indices (s1, s2 & st, cf. section 4.2) a model with 7 factors requires to execute the model at least 16000 times, taking n = 1000. whether the assumption for n was chosen sufficiently large can be judged from the confidence intervals. this results in over 200.000 model evaluations in total due to a relatively finely chosen incremental bending moment rise. in a general case, the validity and robustness of a composite indicator, in this case the reduction factor kc may depend on a number of factors. • the model chosen for estimating the measurement error in the data • the mechanism for including or excluding indicators in the index; the choice of factors fed into a gsa is subjective. particularly in the case of comparatively complex, non-linear numerical models, the model character itself has a considerable influence on the response surface y. theoretically, one could provide the gsa with as many input parameters as possible, but this seems unnecessarily cpu-intensive. not pursued further here, but possible in the first place, seems to be the gradual reduction of the input parameters with subsequent evaluation of the correlation coefficient to the results of a fully packed analysis without static default values. alternatively, qualitative screening using the morris method [21] can be performed as the first step in model survey. • the indicators preliminary treatment; in the investigations presented here, the limiting value 0.65 according to equation 4 is used in the framework of the iteration. the gsa evaluation deliberately excludes this criterion in order to take into account possible influences caused by the asymptotic character of the equation at low principal tensile strains. • the type of normalization scheme applied to the indicators to remove scale effects • the amount of missing data and the imputation algorithm; under high bending moment, it is theoretically possible that the iteration of the strain plane does not reach a stable equilibrium according to the defined stress-strain relationships and limit strains. this leads to a lack of the input value εx,mn in the further iterative consideration of the conditions for the web element, and to a non-response bias depending on the error handling. this scenario was prevented by previous comparative calculations so to avoid inconvergent result data falsifying the gsa. 4.2. sobol’s method after running the numerical model for sobol’s sequence, one gets an array that forms the so-called response surface. sobol’s method evaluates the part of the total variance of the response that can be attributed to input parameter xn. the numerical implementation is based on the salib package [22]. three measures can be obtained for each parameter: • first order index s1: contribution (without interaction) of a parameter to the response variance • second order index s2: interaction of input parameters to the response variance 248 vol. 36/2022 influence of plastic chord deformation on shear capacity • total effect index st: total contribution (including interaction) of a parameter to the response variance a measure of sensitivity is to calculate the variance of the conditional expectation v ar [e(y |xn)] referred to the total variance of y, v ar(y ), cf. equation 6. in other words, the total variance of the response y that can be attributed to input parameter xn defines its weight within the numerical model and, in a further step, enables a better understanding of model-inherent features beyond qualitative screening and, ultimately, opens up initial starting points if a model modification appears desirable. s1n = v ar [e(y |xn = xn)] v ar(y ) (6) any numerical model is some sort of mapping input parameters to output results. the following equation 7 shows the added parts that influence the result y, namely scalar values, the sum of functions evaluated for each parameter (modelling the effect of each individual parameter) and the interaction of parameters. y = f (x) =f0 + np∑ n=1p fn (xn) + ∑ 1≤np fn,n′ (xn, xn′ ) + . . . + f1,2,...,np ( x1, . . . , xnp ) (7) the total variance of the response y (the total of all y-evaluations) can be decomposed into partial variances, attributing variability of the response y to each input parameter, including interactions. different conditional variances: dn = v ar [e (y |xn)] (8) variance on two conditions: dn,n′ = v ar [e (y |xn, xn′ )] − dn − dn′ (9) the first order sobol’ index s1n calculates the impact of the input parameter xn by estimating the partial variance of y explained by this parameter. it estimates by how much the variance of the response is reduced, on average, when the parameter xn is fixed, i.e. it measures the contribution of the parameter xn to the total variance of the response. the total effect index for a parameter xn is defined: stn = sn + ∑ n≤n′ sn,n′ + . . . (10) the total effect index represents the total contribution (including interactions) of a parameter xn to the response variance; it is obtained by summing all first-order and higher-order effects involving the parameter xn. further input variables of the model that are not included in the gsa, but are nevertheless to explicitly assume a variable character, require a recalculation of the sensitivity indices for all sobol’ samples. in the present case, this is the case for the applied bending moment from external load. this procedure is comparatively computationally intensive, but it allows a view on a potentially variable prioritisation of the parameters according to table 1, which should be quite common, especially for non-linear model considerations. 5. results and discussion the gsa shows which of the selected input parameters have an influence on kc and which values would have been sufficiently taken into account within the scope of the investigations with the assumption of a constant value, since no significant contribution to the variance of the result is evident. the major factor is the transverse reinforcement ratio. it is the most important influencing variable over the complete variation of the moment acting on the cross-section. this seems to be plausible with regard to equation 2. described from a phenomenological point of view, the stirrup strain imposes a proportional transverse strain on the concrete compression struts, which reduces the effective concrete compressive strength. for low shear reinforcement levels, this issue is particularly critical because the small cross-sectional area mobilises high strains early on for comparatively small forces after shear cracking. depending on the crack opening and ductility of the reinforcement, rupture of the reinforcement may occur. further questions arise in this regard, which will not be discussed in detail here. a model without parameter interaction leads to∑ st = 1, which is obviously not the case here. under external bending moment of about 0.45 to 0.5 mnm, interesting dependencies between longitudinal and shear reinforcement ratio are revealed, cf. figure 5. the concrete compressive strength also appears to gain in importance in the meantime before st settles back down to the initial level. figure 6 shows also a remarkable interaction between ρsw and εc2 develop in this range. this intermediate effect is due to the limiting conditions of the possible bandwidth of permitted compression strut angles. after exceeding the decompression stresses, small longitudinal strains of the same order of magnitude as εc2 initially enter into the equations. as expected, the influence of the longitudinal strain rises under increasing moment, while the shear reinforcement ratio, which is still important, decreases significantly. a moment of 0.9 mnm for the considered example cross-section (figure 2) results in the longitudinal reinforcement reaching its yield stress. the high longitudinal strain εx,mnv becomes more important for the iterative determination of the compression strut angle and the principal strain ε1 under increasing load. against the background of high utilization rates for bending and shear of many existing 249 sebastian thoma, oliver fischer acta polytechnica ctu proceedings my [mnm] asl asw asl asl asw figure 5. individual parameter importance s1 and total prioritisation st affecting kc concrete bridges, the question arises to what extend the mostly constant factors in standards provided represent a reasonable assumption. directly linked to this is, of course, the question of an adequate formulation of the effective concrete compressive strength or, at a distance, the question of the fundamental capacity of the model conception for shear strength discussed here. the generally constant assumption of the principal strain εc2 even proves to be a definite parameter to be taken into account, largely independent of the acting moment, if one considers modifying the numerical model with respect to kc. the importance of a reasonable assumption or parameterized formulation for εc2 becomes clear when looking at figure 3. selected principal compressive strain and longitudinal strain εx are of the same order of magnitude in absolute terms and have an influence on the compression strut angle according to their difference. 6. conclusions the paper presents a numerical model that allows a closed-loop determination of the effective concrete compressive strength in a beam web under shear loading. of further interest is the reduction factor kc used, which is influenced to different degrees by different parameters. in order to determine the individual influence of the parameters and possible model-determined interaction effects, a gsa was carried out, which was able to identify the shear reinforcement ratio and the assumption of the principal strain εc2 in general, and my [mnm] s2 figure 6. s2 indices: input parameter interaction affecting kc the longitudinal reinforcement ratio and corresponding, ultimately plastic chord deformation under high moment loading in particular as the decisive adjusting screws. in light of the experimental investigations, which achieved high shear capacities despite plastic chord deformation, a few questions arise: • do the empirically derived formulations for determining the reduction factor kc sufficiently approximate the conditions in a prestressed beam web or are the model ideas of plasticity theory described here (especially for cross-sectional segmentation) not applicable? • which possibilities for a more refined estimation of the main strain εc2 seem practicable and reasonable? these questions are part of further considerations, which should be evaluated in the context of the ductility of the reinforcement, which was already briefly addressed in the course of the discussion of results. list of symbols kc reduction factor for concrete compressive strength due to transverse tensile strain [–] εc concrete strain [–] εx,mn axial concrete strain due to normal forces and bending moment at half the height of the inner lever arm [–] εx,v axial concrete strain due to shear [–] ε1 concrete principal tensile strain [–] 250 vol. 36/2022 influence of plastic chord deformation on shear capacity ε2 concrete principal compressive strain [–] θ compression strut angle [deg] fce effective concrete compressive strength [mpa] fc concrete compressive strength, derived from cylinder tests [mpa] ρsl longitudinal reinforcement ratio, defined as ρsl = asl/(bw · d) [–] ρsw geometric shear reinforcement ratio, defined as ρsw = asw/bw [–] λ shear slenderness, defines as λ = m/(v · d) [–] zm lever arm of inner forces [mm] zs lever arm between fc and fs [mm] zp lever arm between fc and fp [mm] σc concrete (compressive) stress [mpa] σs steel stress due to normal forces and bending moment [mpa] σp prestress in tendons [mpa] ∆σp increased tendon stress due to normal forces and bending moment [mpa] αp tendon slope [deg] f c resultant inner compressive force [mn] f t resultant inner tension force as weighted sum of fs and fp [mn] v shear force [mn] es modulus of elasticity of longitudinal reinforcement [mpa] ep modulus of elasticity of prestress tendon strands [mpa] as cross-sectional area of longitudinal reinforcement in tension [m2] ap cross-sectional area of prestressed tendons [m2] s1 sobol’s first order index individual parameter contribution to total response variance [–] s2 sobol’s second order index contribution of parameter interactions (higher-order effects) to total response variance [–] st sobol’s total index combines margins of s1 and s2 for each parameter [–] references [1] n. schramm, o. fischer. zur anrechenbarkeit von nicht normgemäßen bügelformen auf die querkrafttragfähigkeit von bestandsbrücken. bauingenieur 95(11):408–418, 2020. https://doi.org/10.37544/0005-6650-2020-11-66. 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[20] cen. eurocode 2: design of concrete structures part 2: concrete bridges design and detailing rules. tech. rep., european committee for standardization, 2010. 251 https://doi.org/10.37544/0005-6650-2020-11-66 https://doi.org/10.1002/suco.201900085 https://doi.org/10.37544/0005-6650-2020-11-55 https://doi.org/10.1002/9780470725184 https://doi.org/10.1002/best.201900022 https://doi.org/10.3929/ethz-a-001734903 https://doi.org/10.14359/907 https://doi.org/10.14359/10416 https://doi.org/10.3929/ethz-a-001945805 https://doi.org/10.35789/fib.bull.0055 https://doi.org/10.14359/16438 https://doi.org/10.1038/s41592-019-0686-2 https://doi.org/10.14359/51682989 https://doi.org/10.1016/s0378-4754(00)00270-6 https://doi.org/10.1016/s0010-4655(02)00280-1 2008.08051 sebastian thoma, oliver fischer acta polytechnica ctu proceedings [21] m. morris. factorial sampling plans for preliminary computational experiments. technometrics 33(2):161–174, 1991. https://doi.org/10.1080/00401706.1991.10484804. [22] j. herman, w. usher. salib: an open-source python library for sensitivity analysis. the journal of open source software 2(9), 2017. https://doi.org/10.21105/joss.00097. 252 https://doi.org/10.1080/00401706.1991.10484804 https://doi.org/10.21105/joss.00097 acta polytechnica ctu proceedings 36:244–252, 2022 1 introduction 2 experimental evidence 3 compression softening approach 3.1 background 3.2 numerical model 4 global sensitivity analysis 4.1 sobol's sequence 4.2 sobol's method 5 results and discussion 6 conclusions list of symbols references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.35.0023 acta polytechnica ctu proceedings 35:23–26, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague railway track deflection analysis by using evolutionary algorithms pavel kulich∗, otto plášek brno university of technology, antonínská 548/1, 601 90 brno, czech republic ∗ corresponding author: pavel.kulich@vutbr.cz abstract. in contrast to numerical methods, analytical modelling of the railway track is one of the less time-consuming and computationally demanding methods which, in combination with the computing power commonly available today, can form an effective tool for analysing the behaviour of the railway track. this paper deals with the use of iterative methods of evolutionary algorithms, together with analytical modelling, for the purpose of reverse analysis of the measured deflection caused by moving loads acting on the railway track. the theoretical assumptions of the analytical model used, the data collection methodology and the method used for the reverse analysis are presented. the results of the analysis are also presented. keywords: railway track deflection, reverse analysis, evolutionary algorithms. 1. introduction the analytical model presented in [1] was used for the reverse analysis. the model consists of two infinitely long beams lying on the pasternak foundation, which, in contrast with the winkler foundation, considers the shear interaction of adjacent elements. graphical model interpretation is shown in figure 1. the model can be interpreted so that the first layer represents the rail; the second represents the sleeper with substructure. the influence of dynamic variables is taken into account in the same way as in the frýba model [2]. the model can be described by the following differential equations: ei1 d4w1(x, t) dx4 + m1 d2w1(x, t) dt2 + c1 dw1(x, t) dt + k1[w1(x, t) − w2(x, t)] = 0, (1) ei2 d4w2(x, t) dx4 − ga dw2(x, t)2 dx2 + m2 d2w2(x, t) dt2 + c2 dw2(x, t) dt + (k1 + k2)w2(x, t) − k1w1(x, t) = 0, (2) where ei1 is upper beam bending stiffness, m1 is the mass of one meter of the upper beam. according to the paragraph above, the upper beam represents the rail, but the values of ei1 and m1 do not necessarily correspond to the mechanical parameters of the rail itself due to the interaction of the rail with other elements of the railway track. the parameters k1 and c1 are parameters that act between the first and second layers of the structure and can be interpreted as properties of the fastening system. the bottom layer figure 1. schematic of the two-layer model on pasternak foundation. of the model corresponds to the sleepers and partly to the substructure with parameters ei2 bending stiffness of the second layer, m2 mass of one meter of the second layer, ga shear stiffness of the bottom layer. sleepers are not capable of transferring bending loads in the longitudinal direction of the track, and therefore the bending stiffness ei2 can be understood as the residual ability of the sleeper layer to transfer (when the axle passes) the bending load by horizontal action on the trackbed. this stiffness, therefore, takes on a non-zero but negligible value. the parameters k2 and c2 represent the properties of the substructure. it is necessary to introduce the relative coordinate s to solve the above differential equations. the effect of the wheel load q is introduced during the solution through boundary conditions. by substituting the parameters ei1 = 4500 kn m2, ei2 = 0.1 kn m2, ga = 6000 kn , k1 = 250000 kn m−2, k2 = 40000 kn m − 2, v = 30 ms − 1, c1 = 90 kn sm − 2, c2 = 120 kn sm − 2, m1 = 60 kg, m2 = 300 kg, q = 100 kn into the solution matrix, we obtain the deflection curve for the upper layer w1 and the deflection curve for the bottom layer w2. 23 https://doi.org/10.14311/app.2022.35.0023 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en pavel kulich, otto plášek acta polytechnica ctu proceedings figure 2. output of the two-layer model. the horizontal axis of both deflection curves is in the dimensionless relative coordinate s. 2. methods 2.1. track vertical displacement measuring method the input data for the reverse analysis are the track vertical displacements measured in-situ. vertical displacement of railway track is measured relatively to a plain of substructure. the displacement sensors were positioned in the monitored section in such a way that it was possible to detect the pushing of the rails and sleepers. the rail foots and the sleepers in four points along longitudinal axis were monitored. the sensors were positioned as follows: • s0 – right sleeper head; • s1, s2 – right rail; • s3, s4 – sleeper in thirds; • s5, s6 – left rail; • s7 – left sleeper head. using this sensor set, it is possible to obtain the deflection value of the rail unloaded by the twisting of the cross-section. the rail deflection is easily calculated as the average of the values measured by the pair of sensors s1, s2 and s5, s6. using sensors placed on the sleeper, the sleeper deflection curve can be calculated for each measured moment using an interpolation polynomial (cubic spline). it is also possible to calculate the deflection value of the sleeper under the rail at a point that is otherwise inaccessible for sensor placement. figure 3. schematic of sensor placement [3]. 2.2. evolutionary algorithm method reverse analysis of measured data can be classified as an optimization problem. the quantities to be optimized are the parameters of the analytical model. when solving the problem, we look for a combination of model parameters that leads to the best possible solution. the quality of the solution is assessed by the so-called fitness function, which is a mathematical expression of the similarity of the measured data and the output of the analytical model for the given parameters. in general, the higher the quality of the solution, the higher the value of the fitness function. for reverse analysis, algorithms have been developed that use as input a set of random combinations of parameters of the analytical model, successively test each set of parameters, evaluate their quality and from the best combinations create (by precisely defined methods) combinations of new parameters, which test repeatedly. this method of evaluation 24 vol. 35/2022 railway track deflection analysis by using evolutionary algorithms figure 4. measured in-situ deflection leads to better quality solutions with each subsequent iteration. the number of iterations and the number of elements of the parameter set are the input parameters of the algorithm. the optimisation is completed when the required number of iterations is reached. 3. model results the optimization algorithms were tested on data measured in planá nad lužnicí, km 74,978 section without under sleeper pads. track superstructure consists of 60 e1 rails, w14 fastening system and b 91s/1 sleepers. assessed data trace is measured during the passage of the passenger train. data traces from sensors s2 and s6 were selected for further processing, and their values were averaged at the corresponding time points. the resulting data trace was used for comparison with the deflection w1 of the analytical model. all axles of the passenger train were considered. the analytical model presented above gives the deflection for only one axle. to obtain a total deflection line for multiple axles is necessary to introduce the superposition principle for several deflection lines. figure 4 shows that the axles exert different forces on the superstructure, which is also taken to account during superposition calculation. the displacement line obtained by the presented analytical model captures the main characteristics of the measured signal, such as the travel wave in front of the first axle, the travel waves between axles and the steepness of the drop of the track under the passing axle. parameters of the analytical model are shown in table 1. the track critical speed is also calculated. 4. conclusion this article discussed reverse analysis theoretical presumptions, data collection methodology, data processing methodology. the algorithm output is also compared with in-situ measured data. from the text above it can be concluded that the presented analytical model can be used for further analyses of the rail track dynamic behaviour. despite the convincing qualitative results, the evolutionary algorithms, because of its stochastic character, can lead to misleading quantitative outputs. examination of the applicability of evolutionary algorithms for the purposes of reverse analysis, or the incorporation of other methods, will be the subject of the author’s further work. list of symbols ei1 upper layer bending stifness [n m2] ei2 bottom layer bending stifness [n m2] ga bottom layer shear stifness [n] k1 fastening system spring stifness [n m−2] k2 foundation spring stifness [n m−2] c1 fastening system damping [ns m−2] c2 foundation damping [ns m−2] m1 upper layer mass [kg m−1] m2 foundation layer mass [kg m−1] q wheel force [n] v vehicle speed [km h−1] vcr critical speed [km h−1] x longitudinal coordinate [m] t time [s] s relative coordinate [–] w1 rail layer deflection [m] w2 sleeper layer deflection [m] 25 pavel kulich, otto plášek acta polytechnica ctu proceedings figure 5. analytical model comparison with measured data ei1 ei2 ga k1 k2 c1 c2 [kn m2] [kn m2] [kn ] [kn m−2] [kn m−2] [kn s m−2] [kn s m−2] 5428 0,739 549 403287 66903 17 190 m1 m2 v vcr [kg m−1] [kg m−1] [km h−1] [km h−1] 292 515 90 1018 table 1. analytical model parameters acknowledgements the paper was prepared within the project "affordable railroad smart sensing system 4.0", project number tm01000016, which is co-financed with the state support of the technology agency of the czech republic within the 1st public competition of the support programme for applied research, experimental development and innovation delta 2 2019. references [1] p. kulich. dynamic analysis of track. diploma thesis, brno university of technology, 2017. [2] c. esveld. modern railway track. mrt productions, zaltbommel, second edi edn., 2001. [3] o. plášek, j. smutný, r. svoboda, m. hruzíková. analysis of railway track dynamic parameters. in evaces 09. experimental vibration analysis for civil engineering structures, pp. 239–240. wroclaw university of technology, wroclaw university of technology, 2009. 26 acta polytechnica ctu proceedings 35:23–26, 2022 1 introduction 2 methods 2.1 track vertical displacement measuring method 2.2 evolutionary algorithm method 3 model results 4 conclusion list of symbols acknowledgements references https://doi.org/10.14311/app.2022.33.0411 acta polytechnica ctu proceedings 33:411–416, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague physical properties of portland cement based concrete exposed at a depth of 3520 m in the nankai trough shun nomuraa, ∗, takafumi kasayab, yuki sakoic, hisashi fukadad, akira matsumotoe a japan agency for marin-earth science and technology, research institute for value-added-information generation, center for mathematical science and advanced technology, 3173-25 showa-machi kanazawa-ku yokohama-city kanagawa, 236-0001, japan b japan agency for marin-earth science and technology, research institute for marine resources utilization, submarine resources research center, 2-15 natsushima-cho yokosuka-city kanagawa, 237-0061, japan c hachinohe institute of technology, faculty of engineering, department of civil engineering and architecture, 88-1, obiraki, myo, hachinohe, aomori, 031-8501, japan d fudo tetra corporation, geo-technical division, 7-2 nihonbashi komai-cho chuou-ku, tokyo, 103-0016, japan e fudo tetra corporation, technical research institute, 2-7 higashi nakanuki, tsuchiura, ibaraki, 300-0006, japan ∗ corresponding author: nomura.shun@jamstec.go.jp abstract. concrete is widely used in large-scale construction of submarine infrastructure because of its high strength, durability, and ease of handling. however, knowledge of its durability in deep seawater is lacking. in the deep sea, materials are exposed to high pressures and low temperatures, which may cause early deterioration of concrete over time. concrete materials may also be affected by the chemical composition of seawater, which induces the leaching of calcium. in situ exposure tests are therefore important for understanding degradation processes in the deep sea. in this study, portland cement based concrete specimens were placed at a depth of 3520 m on the northern edge of the nankai trough in 2018 and retrieved in 2019, in the deepest exposure testing conducted to date. here we provide an outline of the tests, describe the physical properties of materials exposed to deep seawater, freshwater, and air, and discuss possible concrete degradation mechanisms. keywords: deep sea, exposing test, physical property. 1. introduction the deep sea is regarded as the ’last frontier’ and has attracted much attention in earth science [1–3]. recent studies have indicated that the deep sea has potential for the production of natural resources and oceanic energy. there is also interest in the monitoring of ocean-bottom dynamics and management of subsea disasters [4]. carbon dioxide capture and storage under the seafloor may lead to further utilization of sea-bottom areas [5]. submarine infrastructure with long-term stability is essential for increasing resource productivity and improving monitoring accuracy. concrete is the most preferred material for infrastructure construction because of its cost-effectiveness and rigidity. however, there may be problems with its use in the deep sea, with existing field data being available for depths of up to only 200 m, and with concerns over the durability of concrete in deep-sea environments. concrete strength may decrease due to the effects of extremely high hydrostatic pressures and low temperatures. chemical reactions in the deep-sea environment must also be considered. the carbonate compensation depth (ccd) of the pacific ocean is of the order of 3000 m. below that depth, calcium-one of the main components of concrete-may be leached from concrete, promoting its degradation. there may also be other factors, which together affect the durability of concrete in the deep sea. however, the field data concerning concrete properties in the deep sea are lacking, and there are concerns in extrapolating from shallow-water data. deep-sea exposure testing was therefore undertaken in this study. here we describe the testing of ordinary portland cement based concrete specimens at a depth of 3520 m in the nankai trough, pacific ocean, in the deepest concrete testing yet undertaken. 2. method research cruises were undertaken in 2018 and 2019 as described below. 411 https://doi.org/10.14311/app.2022.33.0411 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en s. nomura, t. kasaya, y. sakoi et al. acta polytechnica ctu proceedings figure 1. field location. figure 2. concept (left) and image (right) of the platform. 2.1. field location the concrete exposure area lies on the northern edge of the nankai trough, off sagami bay, japan, at 34◦1.65’n 138◦31.50’e (figure 1). multi-beam echo-sounder (mbes), bathymetric, and bottom-imaging data were obtained by the research vessel ’yokosuka’ at locations satisfying the following conditions: depth > 3000 m (below the ccd); ease of access (within a two-day cruise), maximum eight-hour dive, solid seabed, and stable tidal systems. 2.2. specimen placement a stainless-steel platform (2.2 × 2.2 × 1.5 m; 600 kg weight) was installed to mount specimens on and to reduce effects of benthic organisms and loss in the strong bottom current (figure 2). the platform was based on four 200-mm diameter discs to prevent sinkage in the sediment. on 18th july 2018, the platform was lowered from the ’yokosuka’ by winch at 45 m.min−1. concrete specimens (36) were placed in six plastic baskets (400×272×300 mm; figure 3). specimen size was 100 mm diameter by 200 mm length (to fit basket pockets). on july 19th, the specimens, in the baskets, were installed on the platform using the deep-sea submersible research vehicle, ’shinkai 6500’, as shown in figure 4. 2.3. retrieval of specimens on 29th september 2019, the study area was accessed by the research vessel ’kairei,’ and the remotely operated vehicle ’kaiko’ was used to retrieve the specimens after 437 days of exposure. the overview of the research cruises in 2018 and 2019 is presented in figure 5. 3. materials and exposure results 3.1. preparation of concrete specimens the concrete mix used for the specimens is described in table 1. ordinary portland cement (density 3.15 g cm−3) was used with a water/cement ratio of 0.55. natural sand (density 2.64 g.cm−3) and crushed limesand (density 2.66 g.cm−3) were mixed and used as a fine aggregate. coarse aggregate was prepared from crushed limestone (5 − 15 mm particles, density 2.67 g.cm−3; 15 − 20 mm particles, density 2.70 g.cm−3). admixtures included an air entraining agent (i.e., 412 vol. 33/2022 physical properties of portland cement figure 3. concept (left) and image (right) of the plastic basket the for concrete specimen. figure 4. installation of the basket by the manipulator (left) and top view of the exposure system (right). figure 5. overview of placement and retrieval processes in 2018 and 2019. 413 s. nomura, t. kasaya, y. sakoi et al. acta polytechnica ctu proceedings unit content [kg.m−3] admixture w/c s/a w c s g ae agent waterreduction agent [%] [%] limesand natural sand 5 − 15 mm 15 − 20 mm [a*] [c × %] 55.0 41.0 157 285 230 532 553 559 3.5 0.05 ∼ 0.2 * 1a = 0.001% table 1. the concrete mix. figure 6. specimens retrieved from the deep sea. deep sea water salty water exposing term 437 days number of specimens 3 2 2 compression strength fc (mpa) 35.4 33.0 39.3 maximum strain ea (1e-3) 1.89 1.62 1.70 young’s modulus es (kn/mm2) 22.1 30.5 35.4 table 2. averaged compression test values. ae) agent and a water reduction agent. the slump and air content of the fresh concrete were 8.5 cm and 4.7%, respectively. the cast specimens were solid cylinders(!100×200 mm) left in the formwork for 1 day before curing in a water bath for 28 days. the water-cured specimens were covered with polypropylene sheet to prevent drying and stored at room conditions pending use. exposure tests were also undertaken in fresh water and salty water under atmospheric pressure. 3.2. description of degraded specimens recovered specimens are shown in figure 6. there were no obvious cracks affecting the overall specimen structure. exposed aggregate was evident on specimen surfaces. there were several spreading white circular spots of 1 − 3 cm diameter. the mass density increased by ∼ 1% during exposure, but specimen dimensions were almost the same as after curing for 28 days. 3.3. compression testing one-dimensional compression test results obtained by a hi-acits-3000 (marui & co. ltd., japan) automatic compression testing apparatus are shown in figure 7, where profiles represent results of deepsea, fresh, and salt water. averaged physical data for the tests are given in table 2. the deep-sea specimen plot has a lower gradient than that of the other cases, indicating a lower elastic modulus. compression strengths ranged from 33.0 to 39.3 mpa with small variations, regardless of exposure state, with the maximum strain at the compression strength being the highest for the deep-sea specimens. the young’s modulus values of the deep-sea specimens were ∼ 67% of that of the specimen exposed to salt water. the relationship between compression strength and young’s modulus is depicted in figure 8 where solid and dotted lines represent the upper and lower criteria for materials used in the construc414 vol. 33/2022 physical properties of portland cement figure 7. results of compression testing. figure 8. relationship between compression strength and young’s modulus. tion of buildings and bridges, among others, set by the architectural institute of japan [6]. all deep-sea specimens plot below the lower limit, meaning these specimens were sufficiently degraded to fall short of architectural criteria. 3.4. travel velocity of ultrasound ultrasound travel velocity, vp, was determined using a pangit pl-200 (fts ltd., japan) instrument, with results as shown in figure 9. the deep-sea specimens exhibited a decrease in vp relative to the other specimens (18% less than that of salty-water specimens). based on elastic wave theory, poisson’s ratio ν is expressed as: ν = 1 − a + ! |9a − 1||a − 1| 4a , where a = v 2p ρ ed (1) where, ρ[kg.m−3] is density, and ed " kg.m−1.s−2 # is the dynamic viscosity coefficient. figure 10 shows the ν values of specimens exposed to fresh and salt water, representing the general values for concrete materials (ν = ∼ 0.25) [7]. in contrast, the lower vp value of deep-sea specimens (ν = ∼ 0.39) is similar to that of non-compressive material (ν = ∼ 0.50). figure 9. ultrasound travel velocity. x-axis represents the exposure conditions and number of specimens. figure 10. poisson’s ratio ν. x-axis represents the exposure conditions and number of specimens. 4. discussion the compression and ultrasonic tests confirm that the mechanical states of concrete specimens exposed in the deep sea were different to those exposed to fresh and salt water, with the deep-sea specimens being more ductile and softer. a conceptual representation of the deep-sea exposure environment is provided in figure 11, illustrating how specimens underwent rapid isotropic loading and unloading (0-35 mpa) during emplacement and retrieval. the specimens were also exposed to high pressures and low temperatures (1−2◦c) after installation. it, therefore, seems that the specimens were degraded by short-term mechanical and long-term chemical actions. to elucidate high loading and unloading effects, cui et al. [7] performed compression tests after exposure of deep-sea samples to high hydrostatic pressures of 30-500 mpa. they confirmed that as hydrostatic pressure increases, the compressive strength decreases, the strain at compressive strength increases, and the young’s modulus decreases. it was also confirmed by scanning electron microscope (sem) observations that, even in an isotropic environment, anisotropy develops inside the concrete specimens due to extremely high pressure, with fine cracks developing in the damaged concrete, which had a poisson’s ratio of almost 0.4. these results are consistent with 415 s. nomura, t. kasaya, y. sakoi et al. acta polytechnica ctu proceedings figure 11. conceptual representation of the physical and chemical environments around in the concrete specimens. those for the deep-sea specimens, suggesting that differences in mechanical history contribute to the deterioration of concrete. 5. conclusion concrete specimens were exposed at a depth of 3520 m for 437 days and their physical properties studied. the specimens deteriorated in the deep-sea environment. in this study, we introduced only those physical properties shown by the compression and ultrasonic tests, whereas the deterioration mechanisms are currently being elucidated by mercury injection tests, sem, electron microprobe analysis, and x-ray diffraction studies. these studies will aid the understanding of physical and chemical processes that affect concrete in deep-sea environments more precisely. further samples are scheduled to be collected over the next five years, and the scientific reasons for their deterioration may become clearer. acknowledgements we thank dr hide sakaguchi and mr masahiko kamei at japan agency for marine-earth science and technology and prof. kenji kaneko of the hachinohe institute of technology for their useful discussion, research assistance, and encouragement. references [1] e. araki, d. m. saffer, a. j. kopf, et al. recurring and triggered slow-slip events near the trench at the nankai trough subduction megathrust. science 356(6343):1157-60, 2017. https://doi.org/10.1126/science.aan3120. [2] r. sharma. environmental issues of deep-sea mining. procedia earth and planetary science 11:204-11, 2015. https://doi.org/10.1016/j.proeps.2015.06.026. [3] k. a. miller, k. f. thompson, p. johnston, et al. an overview of seabed mining including the current state of development, environmental impacts, and knowledge gaps. frontiers in marine science 4, 2018. https://doi.org/10.3389/fmars.2017.00418. [4] w. o. mccarron. deepwater foundations and pipeline geomechanics, florida: j. ross publishing, pp. 336, 2011. [5] s. m. benson, f. m. orr. carbon dioxide capture and storage. mrs bulletin 33(4):303-5, 2011. https://doi.org/10.1557/mrs2008.63. [6] reinforced concrete structure calculation standard and explanation, 9th edition, (tokyo: architectural institute of japan). [7] j. cui, h. hao, y. shi, et al. experimental study of concrete damage under high hydrostatic pressure. cement and concrete research 100:140-52, 2017. https: //doi.org/10.1016/j.cemconres.2017.06.005. 416 https://doi.org/10.1126/science.aan3120 https://doi.org/10.1016/j.proeps.2015.06.026 https://doi.org/10.3389/fmars.2017.00418 https://doi.org/10.1557/mrs2008.63 https://doi.org/10.1016/j.cemconres.2017.06.005 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0483 acta polytechnica ctu proceedings 38:483–487, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague specifying boundary conditions for the operation of pipe heating systems with impact on the building energy balance jakub spurný∗, michal kabrhel czech technical university in prague, faculty of civil engineering, department of indoor environmental and building services engineering, jugoslávských partyzánů 1580/3, 160 00 prague 6 – dejvice, czech republic ∗ corresponding author: jakub.spurny@fsv.cvut.cz abstract. unaddressed boundary conditions in the design of heating systems affect the energy balance of buildings, especially in buildings with very low energy consumption. buildings with very low energy consumption are very sensitive to any heat flow and neglecting realistic heating water parameters affects their energy balance. two simulation models of the end part of the heating system have been developed. the first simulation model shows the effect of the inaccuracy of the designed heating element on the calculated room temperature. the second model shows the hydraulic behaviour of the connection pipes of the designed heating element. keywords: radiator, heating system, energy balance, low energy building heating system. 1. introduction in buildings with low energy consumption, we expect high efficiency of energy systems [1]. the use of hot water systems is most common due to the worst environmental impacts of direct electricity use. therefore, it is important not only to fit new efficient heat sources and use modern control valves but also to focus on the most accurate calculations for their correct design. there is a lot of simplification of individual boundary conditions in the usual heating system design procedures, as they were fully suitable for older and simpler systems. 2. heating system effect the output of the heating elements that provide the indoor ambient temperature should be the same as the heat loss of the heated space. in this case, the heating system operates at its highest efficiency. there are several reasons why this condition may not be met. these include incorrect determination of room heat loss, oversized heating element performance, neglect of heat gain from visible or built-in ductwork, or incorrect heating water parameters – temperature or flow rate. at the same time, a different internal room temperature can also cause heat flows within an energy-efficient building that did not occur previously due to temperature equality. the dependence of the heating performance on the indoor temperature can be determined by calculation. 3. model of the panel radiator the room temperature indicates the calculated internal room temperature and is based on the rules specified in the csn en 12831-1 standard [2]. the standard considers radiators to have the same air temperature as the room temperature. the calculation is solved for steady state in accordance with the standard, and therefore the accumulation does not affect the resulting parameters. heat losses and heat gains are solved in the calculation using their own control elements. the results of the behaviour of control elements are always considered after stabilization of the changed parameters. the following iteration formulas (1), (2) are used to calculate the operating room temperature. t iop = mop · c · (t pop − t zop) hop + t e. (1) t zop = t pop − mn · (t pn − t zn) mop · ( t pop +t zop 2 − msk ·c·(t pop −t zop ) hop − t e t pn +t zn 2 − t in )n . (2) subscripts are used in equations (1) and (2): op operating, n nominal. these subscripts are used for the physical parameters: q radiator heating power [w], m mass flow [kg s−1], c specific heat capacity water [j kg−1k−1], n heating element exponent [-], t p supply water temperature [°c], t z return water temperature [°c], t i room temperature [°c], t e external temperature [°c], h specific heat flux through building structures [w k−1]. 483 https://doi.org/10.14311/app.2022.38.0483 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en jakub spurný, michal kabrhel acta polytechnica ctu proceedings var. 1 var. 2 var. 3 var. 4 var. 5 var. 6 var. 7 var. 8 supply water temperature t pn [°c] 75 75 75 75 55 55 55 55 return water temperature t zn [°c] 65 65 60 60 44 45 40 40 real mass flow mop [kg s−1] = mn mp mn mp mn mp mn mp table 1. boundary conditions to calculate the internal temperature of the room (mn = mass flow for requested power output, mp = mass flow for room heat loss). figure 1. calculated value qn [%] and t iop [°c]. the panel radiator model contains 8 variants of different heating water temperatures and the resulting mass flow rates (see in table 1). design room temperature t in = 20 °c, external temperature t e = −15 °c, heating element exponent n = 1.3 (typical value for flat radiator), room heat loss qp = 350 w. the actual room temperature t iop [°c] and the actual output of the heating body qn [%] were calculated. 3.1. theoretical case study of the panel radiator the results of modelling the given variants are presented in graphical form in the following figure for clarity. from the results in figure 1, it can be seen that the variants with real mass flow mop = mn (nominal value) have larger room temperature deviations t iop than the variants with real mass flow mop = mp when the same % deviation of qn from qp is specified. hence, it appears to be more advantageous to regulate the balancing at mop = mp, as the effect of underrated qn is compensated here, and conversely, overrated qn does not overheat as much again. it can also be seen that the larger deviations of the room temperature t iop are at qn < 100 % than at qn > 100 %, hence more problematic when undersizing the panel radiator. however, in general the variations in room temperature are quite significant for individual qn <,> 100 % with respect to thermal comfort of the occupants in the heated room and the control range of the local heating element control. for qn = 110–120 % the pressure (according to variants t pn and t zn) is out of the range of the thermostatic valve with head. this implies that the valve closes already by design error and is therefore no longer able to react to random heat gains, which must thus legitimately start to overheat the heated space. for the calculated underpressure at qn < 100 % (according to the variants t pn and t zn) it is important whether the valve on the heating element has the possibility of a so-called overflow compared to the balanced kv value e. g. at xp = 2 k and whether this overflow is able to raise the underpressure to the desired state. however, even in this case, the overflow when opening the valve reduces the effectiveness of the subsequent control of the random heat gains. 4. model of the radiator connection pipe today, however, the requirements for systems are more complicated with the growth of many variant solutions, and at the same time computer programs no longer pose obstacles to more demanding mathematical models. one of the boundary conditions addressed is the effect of heating water temperature on the pressure loss in piping systems commonly known calculation for the total pressure loss is the sum of the pressure loss through the inserted resistances and the frictional 484 vol. 38/2022 specifying boundary conditions for the operation of pipe heating . . . tm [°c] ρ [kg/m3] ρ [%] ν [m2/s] ν [%] 75 974,97 100 3,75e-07 100 70 977,92 100,3 4,03e-07 107,4 65 980,69 100,6 4,34e-07 115,7 60 983,28 100,9 4,69e-07 125,0 55 985,69 101,1 5,08e-07 135,5 50 987,92 101,3 5,53e-07 147,4 45 989,97 101,5 6,04e-07 161,0 40 991,84 101,7 6,63e-07 176,6 35 993,53 101,9 7,31e-07 194,7 30 995,04 102,1 8,10e-07 215,9 table 2. comparison of kinematic viscosity ν and water density ρ as a function of mean water temperature tm. pressure loss according to the darcy-weisbach equation (3) [3–5]. table 2 shows a comparison of the deviations of the kinematic viscosity ν and density ρ as a function of the mean heating water temperature tm. from table 2 it can be seen that the influence of ρ is very small within units of %, but the influence of ν is significant. heating systems often operate at variable water temperatures during operation, whether it is the designed temperature gradient, the effect of qualitative regulation [5] or, for example, the temperature drop due to cooling of the heating water in the pipe route [6]. therefore, a comparison of the pressure losses at heating water temperatures from 75 °c to 30 °c at a step of 5 °c was made. other variable inputs for the determination of the pressure loss are the internal pipe diameter d, the wall roughness k and the mass flow rate m with respect to the type of flow. in the following section it is shown how ν and ρ affect the pressure drop through local resistances z and the frictional pressure drop r in the different flow types. pressure loss through the local resistances z depending on the heating water temperature depends only inversely on ρ according to table 2. thus, at tm = 30 °c it is about 98 % compared to tm = 75 °c. for this reason, we will only consider the frictional pressure drop r (3) in the following, where the problem is more complex. r = λ · w2 · ρ d · 2 . (3) in this paper, the flow is considered for analysis in the laminar region according to poiseuille (4). λlam = 64 re . (4) substituting (4) into (3) we get the modified formula for the basic parameters (5). r = 0.04 · m · ν π · d4 . (5) and in the smooth region of flow according to blasius (6). λsmooth = 0.3164 re0.25 . (6) substituting (6) into (3) we get the modified formula for the basic parameters (7). r = 0.0005 · m 74 · ν 14 π 7 4 · ρ 3 4 · d3 , (7) where: r friction pressure loss [pa m−1], m mass flow [kg s−1], d piping diameter [m], ρ water density [kg m−3], λ friction coefficient [-], ν kinematic viscosity [m2 s−1], re reynolds number [-], w water speed [m s−1]. 4.1. theoretical case study of the radiator connection pipe the following example shows the flow in the laminar region for a cu 15 × 1 pipe up to mass flow m = 31 kg s−1 for variable water temperatures. the connection of a heating element with a power of 550 w at ∆t = 15 °c and 350 w at ∆t = 10 °c is solved. for larger pipe dimensions where mass flows are also larger, laminar flow does not occur. as can be seen from figure 2, the deviation increases at tm = 30 °c up to 216 % compared to tm = 75 °c, which is a lot, but it is necessary to look at the absolute values of the pressure drop. these values are at extremes of r = 4.5–9.6 pa m−1 for m = 30 kg s−1, which are very low for heating systems in energy efficient buildings and the difference does not have a significant effect on the resulting hydraulics. in figure 3 we can see that in the smooth flow zone, the change of friction loss depending on the temperature of heating water is caused by both ν and ρ with corresponding exponents. the deviation of friction loss at tm = 30 °c is 119 % in comparison with tm = 75 °c, which is not negligible. this behaviour 485 jakub spurný, michal kabrhel acta polytechnica ctu proceedings figure 2. relative pressure drop r for different mean heating element temperatures tm (laminar flow). figure 3. relative pressure drop r for different mean heating element temperatures tm (smooth zone). of the deviance in the smooth zone is according to blasius constant for all m, before the flow transforms from smooth to transition zone. in general, for the solved variants with boundary conditions of conventional heating systems, the transition from laminar flow to turbulent flow behaves hydraulically smooth at first, even for values with piping roughness k = 0.2 mm. only by further increasing mass flow does the smooth region variant change to a transition region. the rough hydraulic region is never reached in the boundary conditions for normal mass flow in the pipeline. 5. conclusion for all of the above reasons, it is recommended that the heat loss is calculated accurately, and the heating surface is designed for this loss correctly. if we avoid problems with an incorrect design of heating surface performance, we will have a heating system ready for operation with effective control efficiency and a positive effect on thermal comfort. the results also showed that it is preferable to balance the design of the heating surfaces to the flow rate given by the design output, as it compensates for the problem of underand overheating compared to heaters balanced to the flow rate given by the design output. it was also observed that the internal temperature deviation increases faster for underbalanced heaters than for balanced heaters. an important piece of information is that oversizing a heater by about 15 % above the requirement will already overheat the room beyond the normal local control range, leading to cycling. furthermore, the effect of the heating water temperature was pointed out. in laminar flow, which occurs in the heating surface connections in modern heating systems, a large percentage deviation of r was shown with decreasing heating water temperature. it was also found that in the smooth region of turbulent flow, the pressure drop deviations with decreasing temperature go up to 119 %. this flow region is most commonly represented in today’s heating system. this implies that for today’s heating systems in modern low energy buildings, which are typically low temperature due to the small required heating element capacities and the requirements of heat sources such as heat pumps or gas condensing boilers, the pressure drops are proportionally higher for the same flow rate than for older buildings with high heating water temperatures. this has implications for the design and power consumption of circulators and the design of control and balancing valves. 486 vol. 38/2022 specifying boundary conditions for the operation of pipe heating . . . acknowledgements this research was supported by grant sgs21/009/ohk1/ 1t/1. references [1] european union. directive (eu) 2018/844 of the european parliament and of the council of 30 may 2018 amending directive 2010/31/eu on the energy performance of buildings and directive 2012/27/eu on energy efficiency. official journal of the european union 61(l156):75–91, 2018. http://data.europa.eu/eli/dir/2018/844/oj [2] standard csn en 12831-1 energy performance of buildings – method for calculation of the design heat load – part 1: space heating load, module m3-3, 2021. [3] a. avci, i. karagoz. a new explicit friction factor formula for laminar, transition and turbulent flows in smooth and rough pipes. european journal of mechanics b/fluids 78:182–187, 2019. https: //doi.org/10.1016/j.euromechflu.2019.07.007 [4] c. a. dorao, t. langeland, m. fernandino. effect of heating profile on the characteristics of pressure drop oscillations. chemical engineering science 158:453–461, 2017. https://doi.org/10.1016/j.ces.2016.10.009 [5] r. petitjean. total hydronic balancing: a handbook for design and troubleshooting of hydronic hvac systems. tour & andersson ab, 1994. isbn 9163026260. [6] j. spurný, m. kabrhel. effect of heat distribution losses and cooling of heating water on design of heating system [in czech]. (vliv tepelných ztrát rozvodů a ochlazování otopné vody na návrh otopné soustavy). vytápění, větrání, instalace 26(1):2–5, 2017. 487 http://data.europa.eu/eli/dir/2018/844/oj https://doi.org/10.1016/j.euromechflu.2019.07.007 https://doi.org/10.1016/j.euromechflu.2019.07.007 https://doi.org/10.1016/j.ces.2016.10.009 acta polytechnica ctu proceedings 38:483–487, 2022 1 introduction 2 heating system effect 3 model of the panel radiator 3.1 theoretical case study of the panel radiator 4 model of the radiator connection pipe 4.1 theoretical case study of the radiator connection pipe 5 conclusion acknowledgements references https://doi.org/10.14311/app.2022.33.0232 acta polytechnica ctu proceedings 33:232–237, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague contribution of fly ash to mortar strength development under steam and internal curing ryo insakoa, yuki miyoshia, phat tan huynha, b, c, yuko ogawad, ∗, kenji kawaid a hiroshima university, graduate school of engineering, 1-4-1 kagamiyama, higashi-hiroshima, hiroshima 739-8527, japan b ho chi minh city university of technology, faculty of civil engineering, 268 ly thuong kiet street, district 10, ho chi minh city, vietnam c vietnam national university ho chi minh city, 2 linh trung ward, thu duc city, ho chi minh city, vietnam d hiroshima university, graduate school of advanced science and engineering, 1-4-1 kagamiyama, higashi-hiroshima, hiroshima 739-8527, japan ∗ corresponding author: ogaway@hiroshima-u.ac.jp abstract. the purpose of this study is to quantitatively evaluate the effects of steam curing and internal curing on contribution of fly ash to strength development of mortar by using cementing efficiency factor (k-value) that represents strength development performance as a binder of fly ash. in addition, the pozzolanic reaction of fly ash was evaluated from the viewpoint of calcium hydroxide consumption by using thermogravimetry and differential thermal analysis as well as the degree of fly ash reaction by using selective dissolution method. the result indicated that steam curing improved early compressive strength and internal curing improved compressive strength and k-value at all ages. also, a linear relationship between the degree of fly ash reaction and the k-value was shown regardless of the age and the replacement ratio of fly ash. keywords: cementing efficiency factor (k-value), fly ash, internal curing, steam curing, strength development of mortar. 1. introduction fly ash, a by-product of coal-fired power generation, has been commonly utilized as a supplementary cementitious material in concrete to obtain advantages such as long-term strength enhancement and densification of the microstructure. although coal-fired power plant slowly has been shut down in western europe concerning the environmental impact and the fly ash is decreasing, a certain amount of fly ash production is still produced and desired to be utilized. however, there is a concern that the early strength decreases due to replacing a part of cement with fly ash, and most of fly ash is currently used as a cement raw material in japan [1]. steam curing is one of the useful methods for improving the early strength of fly ash concrete, and some properties of steam-cured fly ash concretes have been investigated. yazici et al. reported that when steam curing was performed, the long-term strength of fly ash concrete was low compared with sealed curing at 20 ◦c [2]. meanwhile, pozzolanic reaction of fly ash could continue when the steam-cured fly ash concrete was cured in water after steam curing [3]. on the other hand, there is an internal curing method as a long-term water supply method. internal curing promotes hydration of cement by supplying water from the inside by using internal curing materials with high water absorption. water-absorbing polymer and artificial lightweight aggregate are used as internal curing materials [4, 5]. in addition to these two conventional internal curing materials, roof-tile waste aggregate has been also investigated as an internal curing material. roof-tile is japanese traditional roofing material and sintered clay at high temperature. roof-tile waste aggregate is produced from the roof-tile in its production due to cracking while sintering at high temperature. sekishu kawara, one of the japanese traditional roof tiles, was firstly used as an internal curing material by suzuki et al. [6], and it was reported that the compressive strength of ultra-high strength concrete was improved and the autogenous shrinkage was reduced by partially replacing the coarse aggregate with the roof-tile waste. it was also reported that the strength of steam cured fly ash concrete was improved by using internal curing material [7]. however, the effects of steam curing as well as internal curing on the reaction of fly ash have not been clearly understood. the purpose of this study is to evaluate the effects of steam curing and internal curing on contribution of fly ash to strength development of mortar by using cementing efficiency factor (k-value). furthermore, 232 https://doi.org/10.14311/app.2022.33.0232 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 fly ash for mortar strength development cement fly ash cao (%) 65.09 3.59 sio2 (%) 20.43 56.74 al2o3 (%) 4.95 28.06 fe2o3 (%) 2.52 5.75 mgo (%) 1.62 0.92 na2o (%) 0.25 0.79 so3 (%) 2.86 1.07 loss on ignition (%) 0.76 2.08 density (g/cm3) 3.14 2.23 blaine fineness (cm2/g) 4560 3530 table 1. chemical compositions and physical properties of cement and fly ash. crushed sand roof-tile waste fine aggregate density (g/cm3) 2.58 2.28 water absorption (%) 2.05 9.75 table 2. density and water absorption of crushed sand and roof-tile waste fine aggregate. unit content (kg/m3) water cement fly ash crushed sand roof-tile waste fine aggregate f0s0 278 696 0 1290 0 f20s0 269 537 134 1290 0 f40s0 260 389 260 1290 0 f0s20 278 696 0 1032 228 f20s20 269 537 134 1032 228 f40s20 260 389 260 1032 228 table 3. mixture proportion. the pozzolanic reaction of fly ash was assessed, and the relationship between the fly ash reaction and the k-value of fly ash was discussed. 2. experimental method 2.1. materials, mixture proportion and curing methods paste and mortar specimens were prepared in this study. high-early-strength portland cement and type ii fly ash were used as cementitious materials. these cement and fly ash met jis r 5210 and jis a 6201, respectively. the chemical compositions and physical properties of cement and fly ash are shown in table 1. the water to binder ratio was 40% by mass, and fine aggregate to mortar ratio was 50% by volume. roof-tile used in this study was derived from the sekishu kawara, which composed of approximately 70% sio2, and 20% al2o3. the density in saturated and surface-dry condition and the water absorption of crushed sand and roof-tile waste fine aggregate are shown in table 2. the fly ash replacement ratios were 0% (f0), 20% (f20), and 40% (f40) by mass. the roof-tile waste fine aggregate was used for mortar specimens as an internal curing material by replacing crushed sand at 0% (s0) and 20% (s20) by volume. the mixture proportions are shown in table 3. two curing methods were applied in this study. one was sealed curing at 20 ◦c at all ages (20 ◦c sealed curing), while the other was steam curing at a maximum temperature of 50 ◦c for 6 hours after casting. the steam-cured specimens were demolded at the age of 1 day and stored at 20 ◦c and 60% relative humidity. 2.2. compressive strength and k-value mortar specimens, which were cast into cylindrical molds of 50-mm diameter and 100-mm height, were used for the compressive strength test. the compressive strength test was conducted according to jis a 1108 at the ages of 1, 7, and 28 days. the equivalent cement-to-water ratio ((c/w )eq ) of the fly-ash mortar was calculated from the compressive strength and the relationship between the cement-to-water ratio (c/w ) and the compressive strength of mortar without fly ash. the k-value was obtained from the (c/w )eq , the c/w and the replacement ratio of fly ash (r) as shown in eq. 1 [8]. k = ! (c/w )eq c/w − 1 " · 1 − r r (1) 233 r. insako, y. miyoshi, p. t. huynh et al. acta polytechnica ctu proceedings 2.3. calcium hydroxide (ch) content and its consumption per unit fly ash content paste specimens cast into 16-ml polypropylene bottles were used for examining the ch content by thermogravimetry and differential thermal analysis. the ch content in the sample was measured at the ages of 1, 7, and 28 days. specimens were cut and samples in the size of 2.5 − 5.0 mm range were obtained. the samples were soaked in acetone to stop hydration and dried in a vacuum desiccator for 24 hours. after that the samples were ground into a powder less than 150 µm before the measurements of the degree of fly ash reaction and the ch content. figure 1. compressive strength of mortar under steam curing. the ch consumption per unit fly ash content can be calculated using the following eq. 2. chconsumption = ch (1 − r) − chf a r (2) • chconsumption: ch consumption per unit fly ash content (%) • ch: ch content in paste without fly ash (%) • chf a: ch content in paste with fly ash (%) 2.4. degree of fly ash reaction the sample used for measuring the degree of fly ash reaction was prepared similarly to that for ch content. the degree of fly ash reaction in paste specimens was measured at 1, 7, and 28 days by using selective dissolution method (sdm). to evaluate the reaction ratio of fly ash, ohsawa et al. proposed a method using 2-m hydrochloric acid (hcl) and 5% sodium carbonate (na2co3) aqueous solution [9]. the procedure of sdm is as follows: a 1-g powder sample was put in a pre-weighed centrifuge tube and weighed precisely, and 30 ml of 2-m hcl solution was added in the tube. then, the tube figure 2. compressive strength of mortar under 20 ◦c-sealed curing. was placed in a water bath at 60 ◦c for 15 min to promote the reaction. the solution was stirred occasionally with a glass rod during this step. the tube was removed from the bath and centrifuged at 4000 rpm (revolutions per minute) for 30 sec to separate solid and liquid phases. the liquid phase was decanted as much as possible without disturbing the solid residue in the bottom. after that, the tube was filled with hot water, centrifuged again and decanted. this operation was repeated 3 times. the tube was then filled with 30 ml of na2co3 solution, placed in a water bath at 80 ◦c for 20 min and stirred continuously. the tube was then centrifuged at 4000 rpm for 1 min. the liquid phase was decanted and the residue was washed with hot water and centrifuged. this washing was repeated 3 times. the tube with residue was dried at 105 ◦c for 24 h and then weighed. the degree of fly ash reaction is calculated using the following eq. 3. α = 1 − x (1 − ig) (1 − ig) k1k2 (3) • α : degree of fly ash reaction • k1 : original fraction of fly ash in ignited base • k2 : residue extracted of a 1-g fly ash • x : residue extracted of the hydrated fly ash in a 1-g hydrated sample • ig : loss on ignition of the hydrated sample • ig′ : loss on ignition of the residue 3. results and discussion 3.1. compressive strength the compressive strengths of mortar specimens under steam curing and 20 ◦c-sealed curing are presented in 234 vol. 33/2022 fly ash for mortar strength development figure 3. k-value of fly ash with age. figure 4. ch content with age. figures 1 and 2, respectively. at the age of 1 day, the compressive strength of steam-cured mortar was 10% to 35% higher than that of 20 ◦c-sealed cured mortar. the roof-tile waste fine aggregate replacement increased the compressive strength of mortar by 15% to 20% at all ages, regardless of curing methods. at the age of 28 days, the mortar specimens with 20% fly ash and 20% roof-tile waste fine aggregate (f20s20) had the same compressive strength as that with 0% fly ash and 0% roof-tile waste fine aggregate (f0) in the case of 20 ◦c-sealed curing. meanwhile, the compressive strength of f20s20 was lower than that of f0 in the case of steam curing. this could indicate that steam curing mainly promoted the initial hydration of cement and slightly promoted the pozzolanic reaction of fly ash which can occur at later ages, while internal curing by roof-tile waste fine aggregate could enhance the cement and fly ash reaction at early ages as well as the later ages. -5 -4 -3 -2 -1 0 1 2 3 4 0 7 14 21 28 c h c on su m pt io n pe r un it fly a sh co nt en t ( % , g / fly a sh -1 00 g) age(days) f20 20ºc-sealed f20steam f40 20ºc-sealed f40steam figure 5. ch consumption per unit fly ash content with age. 3.2. cementing efficiency factor (k-value) figure 3 shows the k-value of fly ash under steam curing and 20 ◦c-sealed curing with time. the steam curing had a lower k-value than the 20 ◦c-sealed curing for all mixture proportions. this result also suggests that steam curing mainly promoted the hydration of cement. in addition, the k-value in the case of 20% roof-tile waste replacement (s20) was higher than that in the case of no roof-tile waste. it indicated that the internal curing by using the roof-tile waste fine aggregate promoted the fly ash reaction, and the fly ash contributed to the strength development of mortar. the f40s20 in both curing conditions had a significant increase in the k-value from 7 to 28 days when compared with the f40. this shows that the internal curing continuously promoted the pozzolanic reaction of fly ash for a long term when the replacement ratio of fly ash was high. 235 r. insako, y. miyoshi, p. t. huynh et al. acta polytechnica ctu proceedings figure 6. degree of fly ash reaction with age. 3.3. calcium hydroxide (ch) content and its consumption per unit fly ash content figure 4 shows the ch content under 20 ◦c-sealed curing and steam curing. at the age of 1 day, the ch content of steam-cured specimens was higher than that of 20 ◦c-sealed cured specimens. it could imply that steam curing accelerated cement hydration and ch was produced more than that under 20 ◦csealed curing. at the age of 7 days, then, the ch content for each mixture proportion was almost the same, and increased up to the age of 28 days slightly. figure 5 shows ch consumption per unit fly ash content by fly ash reaction under 20 ◦c-sealed curing and steam curing. at the age of 28 days, the ch consumption per unit fly ash content in the case of f20 with 20 ◦c-sealed curing was greater than that in the case of f20 with steam curing. besides, the unit ch consumption at the ages of 1 day and 7 days was negative. it may indicate that the cement hydration could be promoted by the incorporation of fly ash and more ch was produced in fly ash cement paste compared to the plain cement paste. 3.4. degree of fly ash reaction figure 6 shows the degree of fly ash reaction under steam curing and 20◦c-sealed curing. at all ages, the degree of fly ash reaction in the case of 20 ◦c-sealed curing was higher than that in the case of steam curing. the longer the material age, the greater the difference between steam curing and 20◦c-sealed curing. it could imply that drying after the end of the steam-supply hindered the fly ash reaction. figure 7 shows the relationship between the k-value and the degree of fly ash reaction. the k-value increases as the degree of fly ash reaction increases. the relationship between the degree of fly ash reaction and the k-value can be represented by one straight line regardless of the age and the replacement ratio of fly ash. figure 7. relationship between k-value and degree of fly ash reaction. 4. conclusions the present study investigated the k-value of fly ash regarding compressive strength development, ch consumption as well as the degree of fly ash reaction in order to evaluate the effects of steam curing and internal curing on contribution of fly ash to strength development of mortar. the following conclusions can be drawn within the limits of this study. steam curing improved early compressive strength of mortar regardless of the fly ash or roof-tile replacement. the 20% roof-tile waste fine aggregate replacement increased the compressive strength by 15% to 20% for both 20 ◦c-sealed curing and steam curing. the internal curing could contribute not only to the cement hydration but also to the fly ash reaction. the k-value in the case of 40% fly ash replacement for both steam and 20 ◦c-sealed curing conditions increased from 7 to 28 days by using roof-tile waste fine aggregate. it was suggested that when the replacement ratio of fly ash is high, the effect of internal curing on the pozzolanic reaction may continue for a long time. the k-value in the case of steam curing was lower than that in the case of the 20 ◦c-sealed curing, and it implies that steam curing significantly promoted the cement hydration whereas it slightly promoted the pozzolanic reaction of fly ash. the relationship between the degree of fly ash reaction and the k-value can be represented by one straight line regardless of the age and the replacement ratio of fly ash. references [1] k. ogawa, h. uchikawa, k. takemoto, et al. the mechanism of the hydration in the system c3s-pozzolana. cement and concrete research 10(5):683-96, 1980. https://doi.org/10.1016/0008-8846(80)90032-0. [2] h. yazici, s. aydin, h. yi ‘giter, et al. effect of steam curing on class c high-volume fly ash concrete mixtures. cement and concrete research 35(6):1122-7, 2005. https: //doi.org/10.1016/j.cemconres.2004.08.011. 236 https://doi.org/10.1016/0008-8846(80)90032-0 https://doi.org/10.1016/j.cemconres.2004.08.011 vol. 33/2022 fly ash for mortar strength development [3] j. tomiyama, y. suda, t. saeki, et al. basic study on strength characteristic of fly ash concrete under steam curing. cement science and concrete technology 66(1):359-66, 2012. https://doi.org/10.14250/cement.66.359. [4] k. yokota, s.-i. igarashi. effects of aggregate particles on spatial distribution of superabsorbent polymers for internal curing. cement science and concrete technology 67(1):179-86, 2013. https://doi.org/10.14250/cement.67.179. [5] d. tachibana, k. kimura, k. naitoh. effect of high-performance artificial lightweight aggregate on several properties of lightweight concrete. doboku gakkai ronbunshu 1994(496):89-98, 1994. https://doi.org/10.2208/jscej.1994.496_89. [6] m. suzuki, i. maruyama, t. kawabata. a study on deformation of ultra high strength concrete containing crushed roof tile aggregate. japan concrete institute 29:651-6, 2007.,. [7] y. ogawa, p. t. bui, k. kawai, et al. effects of porous ceramic roof tile waste aggregate on strength development and carbonation resistance of steam-cured fly ash concrete. construction and building materials 236, 2020. https: //doi.org/10.1016/j.conbuildmat.2019.117462. [8] i. a. smith. the design of fly-ash concretes. proceedings of the institution of civil engineers 36(4):769-90, 1967. https://doi.org/10.1680/iicep.1967.8472. [9] s. ohsawa, e. sakai, m. daimon. reaction ratio of fly ash in the hydration of fly ash-cement system. cement science and concrete technology 53:96-101, 1999. 237 https://doi.org/10.14250/cement.66.359 https://doi.org/10.14250/cement.67.179 https://doi.org/10.2208/jscej.1994.496_89 https://doi.org/10.1016/j.conbuildmat.2019.117462 https://doi.org/10.1680/iicep.1967.8472 https://doi.org/10.14311/app.2022.33.0424 acta polytechnica ctu proceedings 33:424–430, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague comparison between concrete stress diagrams according to the brazilian standard abnt nbr 6118: 2014 victor hugo dalosto de oliveiraa, ∗, rodolfo alves carvalhoa, luana ferreira borgesa, paulo chaves de rezende martinsa, rodolfo rabelo nevesb, joão paulo de almeida siqueiraa, gabriel yves da silva oliveirac, cintia adriana azevedo de liz anhaiad a department of civil and environmental engineering, university of brasília, sg-12, campus darcy ribeiro, asa norte, brasília, df, 70910-900, brazil b department of materials engineering and construction, federal university of minas gerais, av. antø̂nio carlos, 6627, pampulha, belo horizonte, mg, 31270-901, brazil c school of exact, architecture and environment, catholic university of brasília, qs 07 lote 01, taguatinga, brasília, df, 71966-700, brazil d department of civil engineering, faculty of pinhais, av. camilo di lellis, 1065, centro, pinhais, pr, 83323-000, brazil ∗ corresponding author: victordalosto@gmail.com abstract. this study aims to compare methods for the determination of concrete properties by means of the stress diagrams present in the brazilian standard abnt nbr 6118: 2014. the area under the stress diagram, the internal reactions, and the application point of the resulting reactions for the parabola-rectangle and rectangular block diagrams are present in order to compare them. deductions and numerical examples were used, and different results were obtained for each formulation. this is due to non-consideration of the relationship between stress and strain in the simplified rectangular block. the rectangular block is applicable only for cases in which the concrete reaches the ultimate strain. these cases are those that concrete crushing determines the section failure in compression with steel yielding in tension (domain 3) or without steel yielding (domain 4). keywords: concrete behavior, parabola-rectangle diagram, rectangular block. 1. introduction exhaustion of the stress capacity of a reinforced concrete element subjected to normal stress can be achieved by crushing the concrete under compression or by an excessive plastic strain of the steel. however, concrete failure is difficult to identify. it is conventional to assume that this failure occurs when the material reaches a maximum compressive strain determined by experimental results [1]. according to american standard aci 318-14 [2], the maximum compressive strain of concrete is around 0.003 to 0.004 under normal conditions, and it may reach 0.008 under special conditions. the brazilian standard abnt nbr 6118: 2014 [3] states a maximum specific strain of 0.0035 for concretes up to class c50 (characteristic compressive strength up to 50 mpa). this brazilian standard adopts an idealized stressstrain diagram to represent concrete behavior, in which the stress distribution takes place according to a parabola-rectangular diagram [3]. this standard allows to replace the parabola-rectangle diagram with a rectangle of equivalent depth, as a simplification of calculation [3]. abnt nbr 6118: 2014 [3] states that the difference in results with both formulations is small, so there is no need to make any adjustment. the rectangular block does not represent the actual stress distribution within the compressed concrete zone, but provides reasonably the same compressive force [4]. however, mendes neto [5] noticed that the design of structures with the rectangular block provides a different and unsafe result compared to the parabolarectangle diagram result. formulations for stress, internal reaction, and the point of application of the resulting reaction are presented in this paper by means of the parabola-rectangle diagram and the rectangular block. solutions are compared for some cases to verify if the use of the rectangular block leads to reasonable and safe results. 424 https://doi.org/10.14311/app.2022.33.0424 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 concrete diagrams (abnt nbr 6118: 2014) figure 1. idealized stress-strain diagram. source: adapted from the abnt nbr 6118: 2014 [3]. 2. materials and numerical methods this section presents the concepts and formulations considered for the comparison between parabolarectangle and rectangular block diagrams. abnt nbr 6118: 2014 [3] uses an idealized stress-strain diagram to represent concrete compression at the ultimate limit state, as shown in figure 1, in which εc is the concrete strain; εc2 is the assumed specific compression strain of concrete at the beginning of the plastic level; εcu is the assumed maximum useable compression strain in the concrete; fck is the characteristic strength of the concrete; fcd is the design strength of the concrete; and σc is the concrete stress [3, 6]. this diagram is similar to that presented in section 7.2.3.1.5 of the fib model code 2010 [7] and section 3.1.7 of the eurocode 2 [8]. equation (1) defines this diagram (figure 1): σc = 0.85fcd ! 1 − " 1 − εc εc2 #n$ (1) in which n = 2 for fck ≤ 50 mpa; and n is given by equation 2 for fck > 50 mpa. n = 1.4 + 23.4 % (90 − fck) & 100 '4 (2) for concretes with fck ≤ 50 mpa: εc2 = 0.002; and εcu = 0.0035. equations 3 and 4 give the values of εc2 and εcu for concretes with fck > 50 mpa. εc2 = 0.002 + 0.000085 (fck − 50) 0.53 (3) εcu = 0.0026 + 0.035 % (90 − fck) & 100 '4 (4) the brazilian standard [3] states that it is possible to replace this diagram with a rectangle of height y = λx (in which x is the effective depth of the neutral axis), and constant stress αc fcd, without loss of quality in results. for concretes with fck ≤ 50 mpa: λ = 0.8, and αc = 0.85. equations 5 and 6 give values of λ and αc for concretes with fck > 50 mpa. thus, the strains and stresses in a rectangular cross-section at the ultimate limit state can be represented as in figure 2. λ = 0.8 − (fck − 50) & 400 (5) αc = 0.85 % 1 − (fck − 50) & 200 ' (6) the area under the stress diagram can be obtained by integrating stresses in the domain of the specific strains. thus, equation 7 is used when strain values at the most compressed fiber are less than εc2, and equation 8 is used if strain values are larger than εc2. a(ε) = 0.85 fcd ( εc 0 ! 1 − " 1 − ε εc2 #n$ dε (7) a(ε) = 0.85 fcd ( εc2 0 ! 1 − " 1 − ε εc2 #n$ dε + 0.85 fcd ( εc εc2 dε (8) considering the bernoulli hypothesis (plane sections remain plane), the longitudinal strain varies proportionally with the distance to the neutral axis [2]. therefore, if the specific strain is known, the neutral axis is determined. thus, as the resultant of internal reactions equals the volume under the stress diagram, its value becomes known after determining the stress and cross-sectional compressed area. the center of gravity of the compression region is obtained by dividing the stress-strain diagram into infinitesimal areas stress-strain diagram into infinitesimal areas σ dε. the point of application of the resultant of normal concrete stresses can be determined by equation 9, in which the integral of the stress area in the strain domain is determined by equation 7 and 8. ε = ( a xda ( a da = ( ε εσdε ( ε σdε (9) 2.1. strain domains conventionally, failure of a reinforced concrete section occurs when the specific strain of concrete or steel (or both) reaches its ultimate value. the standard abnt nbr 6118: 2014 [3] defines that the ultimate limit state is characterized when the distribution of strains in the section is within one of the specified strain domains. figure 3 illustrates these strain domains, in which εyd is the strain in the tension reinforcement at failure; h is the overall height of a cross-section; d is the distance from the most compressed fiber to the centroid of the longitudinal reinforcement on the tension side of the member; and d′ is the distance from the most compressed fiber to the centroid of the longitudinal compression steel [3, 6]. thus, section failure can occur under excessive plastic strains of the steel characterized by the straight-line 425 v. h. d. d. oliveira, r. a. carvalho, l. f. borges et al. acta polytechnica ctu proceedings figure 2. strain and stress diagrams of concrete. figure 3. strain domains in the ultimate limit state. source: adapted from the abnt nbr 6118: 2014 [3]. a, within domains 1 and 2, or by crushing of the compressed concrete defined by domains 3, 4, 4a, 5, and the straight-line b. the ultimate limit state of domain 2 is characterized by the strain εyd = 1%, and the failure of the cross-section occurs without the concrete reaching its maximum strength. fusco (1981) [1] divides domain 2 into two subdomains: 2a (εc from zero up to εc2) and 2b (from εc2 to εcu). this subdivision aims to identify the value at which the use of compression reinforcement becomes efficient [1]. such consideration becomes important here because it marks the point at which pseudo-plastification of concrete begins. while the section is in subdomain 2a, concrete stresses are less than 0.85 fcd. when concrete strain exceeds εc2, the concrete pseudo-plastification and the constant compression region arise. when the section deformation approaches domain 3, the specific strains become higher, and the plasticized zone expands. therefore, two different equations are required to represent the behavior of compressed concrete, equations 7 and 8. the strain εcu characterizes the ultimate limit state of domains 3, 4 and 4a. in these domains, the failure of the section occurs by crushing the concrete in compression. although the neutral axis goes down in the section and the compressed area increases, the stress diagram proportions remain constant. in these domains, it is possible to find the stresses and reactions in concrete only with equation 8. the strain εcu also characterizes the ultimate limit state of domain 5, with concrete failure in compression. in domain 5, the entire section is under nonuniform compression. the strains converge to a constant value εc2 as the section approaches the line b. thus, the concrete properties in domain 5 are defined by equation 8, changing the lower limit of the first integral. 2.2. calculations with a computer program the matlab software was used for the preparation of scripts in order to make a comparison between the diagrams. the stresses, internal reactions, and center of gravity of the parabola-rectangle diagram were obtained through these scripts, that solve equations 7, 8, and 9. the results were compared with those obtained by the rectangular block. this analysis is called "approach 1". another comparison between the diagrams, with a different approach (approach 2), was performed. values were assigned to the position of the neutral axis, and the internal reactions were obtained with equation 10 for the parabola-rectangle diagram, and with equation 11 for the rectangular block. rc = ( a σ da (10) rc = αc fcd bw λ x (11) 426 vol. 33/2022 concrete diagrams (abnt nbr 6118: 2014) figure 4. concrete strains and stresses in the ultimate limit state. figure 5. representation of calculation solution. 3. results and discussions initially, it should be emphasized that the use of stress diagrams does not apply to line a and domain 1, since there are no compressive stresses in concrete, and its tensile stresses are neglected at the ultimate limit state. 3.1. approach 1 based on the parameters analyzed, the rectangular diagram presents only one way to describe the behavior of the concrete. however, the parabola-rectangle diagram has five different behaviors: 1. while the structure is in domain 2a, there is only a parabolic excerpt (σc < 0.85 fcd); 2. when the pseudo-plastification begins in domain 2b, the excerpt of constant stress 0.85 fcd appears, and there is an expansion of this plastic zone until the section reaches domain 3; 3. for domains 3, 4 and 4a, there is an expansion of the compressed region, in which the stress distribution increases proportionally in the parabolic excerpt and in the constant stress excerpt; 4. for sections in domain 5, there is an overlap of the plastic zone, where the specific strains across the section converge to εc2 and the stresses converge to 0.85 fcd; 5. for line b, the entire section is under uniform compression (σc = 0.85 fcd). figure 4 illustrates the stress distribution and behaviors in the strain domains. results of the parabola-rectangle diagram can be obtained by equations 7 and 8, calculating the area as a function of the neutral axis position (x). for comparison, the solution is represented as proportions of equivalent rectangles of constant stress of 0.85 fcd and nominal height y1 and y2, referring to the portions of constant and parabolic stresses distribution, respectively. figure 5 illustrates the solution, and the results are listed in table 1. domain 4a represents a small region of the domains, and it ends when the section is fully compressed, so its relationships are similar to those of domains 3 and 4. for domain 5 and line b, the neutral axis relation is no longer represented as a function of x/d and becomes a function of x/h, as shown in table 2. thus, the nominal compression of the parabolarectangle stress diagram is similar to the rectangular diagram in domains 3, 4 and 4a, where the height y = 0.8095x is equivalent to the height adopted by simplification (y = 0.80x). however, the point of application of the resulting reaction is different in each type of diagram. in addition, a considerable differ427 v. h. d. d. oliveira, r. a. carvalho, l. f. borges et al. acta polytechnica ctu proceedings (x/d) y/x y1/x y2/x c.g./x domain 2a 0.01 1460 / 29403 ∼= 0.050 1460 / 29403 391 / 1168 ∼= 0.335 0.02 710 / 7203 ∼= 0.099 710 / 7203 191 / 568 ∼= 0.336 0.03 1380 / 9409 ∼= 0.147 1380 / 9409 373 / 1104 ∼= 0.338 0.04 335 / 1728 ∼= 0.194 335 / 1728 91 / 268 ∼= 0.340 0.05 260 / 1083 ∼= 0.240 260 / 1083 71 / 208 ∼= 0.341 0.06 630 / 2209 ∼= 0.285 630 / 2209 173 / 504 ∼= 0.343 0.07 8540 / 25947 ∼= 0.329 8540 / 25947 337 / 976 ∼= 0.345 0.08 590 / 1587 ∼= 0.372 590 / 1587 41 / 118 ∼= 0.347 0.09 3420 / 8281 ∼= 0.413 3420 / 8281 319 / 912 ∼= 0.350 0.10 110 / 243 ∼= 0.453 110 / 243 31 / 88 ∼= 0.352 0.11 11660 / 23763 ∼= 0.491 11660 / 23763 301 / 848 ∼= 0.355 0.12 255 / 484 ∼= 0.527 255 / 484 73 / 204 ∼= 0.358 0.13 12740 / 22707 ∼= 0.561 12740 / 22707 283 / 784 ∼= 0.361 0.14 3290 / 5547 ∼= 0.593 3290 / 5547 137 / 376 ∼= 0.364 0.15 180 / 289 ∼= 0.623 180 / 289 53 / 144 ∼= 0.368 0.16 860 / 1323 ∼= 0.650 860 / 1323 16 / 43 ∼= 0.372 0.166667 2 / 3 ∼= 0.667 2 / 3 3 / 8 ∼= 0.375 domain 2b 0.17 172 / 255 ∼= 0.675 2 / 85 166 / 255 22019 / 58480 ∼= 0.377 0.18 94 / 135 ∼= 0.696 4 / 45 82 / 135 6451 / 16920 ∼= 0.381 0.19 68 / 95 ∼= 0.716 14 / 95 54 / 95 9977 / 25840 ∼= 0.386 0.20 11 / 15 ∼= 0.733 1 / 5 8 / 15 43 / 110 ∼= 0.391 0.21 236 / 315 ∼= 0.749 26 / 105 158 / 315 39211 / 99120 ∼= 0.396 0.22 42 / 55 ∼= 0.764 16 / 55 26 / 55 3697 / 9240 ∼= 0.400 0.23 268 / 345 ∼= 0.777 38 / 115 154 / 345 49859 / 123280 ∼= 0.404 0.24 71 / 90 ∼= 0.789 11 / 30 19 / 45 3481 / 8520 ∼= 0.409 0.25 4 / 5 ∼= 0.800 2 / 5 2 / 5 33 / 80 ∼= 0.412 domains 3 and 4 0.259259 17 / 21 ∼= 0.810 3 / 7 8 / 21 99 / 238 ∼= 0.416∼ 1 table 1. section properties in the ultimate limit state (domains 2, 3 and 4). ence is noted for the nominal compression and the point of application of the resultant reaction within sections in domains 2 and 5, and in line b. 3.2. approach 2 a rectangular section with bw = 1, d = 1, h = 1.2 and fck = 20 mpa was analyzed, in which bw is the section width. regardless of the values adopted for the cross-section, the difference obtained by the diagrams remains proportional. for this reason, dimensionless values were used for bw, h and d. the results for the ultimate limit state are shown in figure 6. as a reference, the limits of the strain domains were marked, where ca-50 steel (steel with 500 mpa yield strength) was adopted. the same fact can be verified for concrete strength group ii (c55 to c90): there is a considerable difference between results with the parabola-rectangle diagram and the rectangular block simplification. a rectangular section with bw = 1, d = 1, h = 1.2 and fck = 80 mpa was analyzed. the relationship between internal reaction and neutral axis position is shown in figure 7. 4. conclusions although the rectangular block simply represents the behavior of the concrete at the imminence of failure, it does not reflect the relation between stresses and strains where concrete is under limit strain of failure. this simplification considers that the material works at its full capacity of strength all over the compressive zone. the nominal compression and point of application of the concrete reaction are almost equal with both diagrams (parabola-rectangle or rectangular block) in the design of linear elements sections submitted predominantly to bending load (domains 3 and 4). however, it can result in unsafe design in sections in domains 2 or 5, typically slabs or columns, respectively. the difference between the diagrams is amplified for high strength concrete elements. since high resistance concretes have lower ductility and an explosive and brittle failure, they have a smaller plastic region and, consequently, values of εc2 close to εcu, with a smaller excerpt of constant stresses. thus, the calculated difference between the stress diagrams becomes larger. 428 vol. 33/2022 concrete diagrams (abnt nbr 6118: 2014) (x/h) y/h y1/h y2/h c.g./h domain 5 1.00 17 / 21 ∼= 0.810 3 / 7 8 / 21 99 / 238 ∼= 0.416 1.05 133349 / 158949 ∼= 0.839 3 / 7 65228 / 158949 805443 / 1866886 ∼= 0.431 1.10 39989 / 46389 ∼= 0.862 3 / 7 20108 / 46389 247923 / 559846 ∼= 0.443 1.15 188621 / 214221 ∼= 0.881 3 / 7 96812 / 214221 1192347 / 2640694 ∼= 0.452 1.20 13709 / 15309 ∼= 0.895 3 / 7 7148 / 15309 87963 / 191926 ∼= 0.458 1.25 10085 / 11109 ∼= 0.908 3 / 7 5324 / 11109 13095 / 28238 ∼= 0.464 1.30 71741 / 78141 ∼= 0.918 3 / 7 38252 / 78141 470187 / 1004374 ∼= 0.468 1.35 323861 / 349461 ∼= 0.927 3 / 7 174092 / 349461 2139027 / 4534054 ∼= 0.472 1.40 5669 / 6069 ∼= 0.934 3 / 7 3068 / 6069 37683 / 79366 ∼= 0.475 1.45 403829 / 429429 ∼= 0.940 3 / 7 219788 / 429429 2698803 / 5653606 ∼= 0.477 1.50 4469 / 4725 ∼= 0.946 3 / 7 2444 / 4725 30003 / 62566 ∼= 0.480 1.75 27725 / 28749 ∼= 0.964 3 / 7 15404 / 28749 37791 / 77630 ∼= 0.487 2.00 2477 / 2541 ∼= 0.975 3 / 7 1388 / 2541 17019 / 34678 ∼= 0.491 3.00 1685 / 1701 ∼= 0.991 3 / 7 956 / 1701 2343 / 4718 ∼= 0.497 4.00 13061 / 13125 ∼= 0.995 3 / 7 7436 / 13125 91107 / 182854 ∼= 0.498 5.00 335 / 336 ∼= 0.997 3 / 7 191 / 336 234 / 469 ∼= 0.499 10.0 94205 / 94269 ∼= 0.999 3 / 7 53804 / 94269 131823 / 263774 ∼= 0.500 50.0 ∼= 1.000 3 / 7 0.571 ∼= 0.500 line b ∞ 1 3 / 7 4 / 7 0.5 table 2. section properties in the ultimate limit state (domain 5 and line b). figure 6. internal reactions for concrete class c20 (fck = 20 mpa). the percentage difference between the two approaches is marked at some representative points, referred to results obtained with the parabola-rectangle diagram. figure 7. internal reaction for concrete class c80 (fck = 80 mpa). the percentage difference between the two approaches is marked at some representative points, referred to results obtained with the parabola-rectangle diagram. 429 v. h. d. d. oliveira, r. a. carvalho, l. f. borges et al. acta polytechnica ctu proceedings acknowledgements this study was financed in part by the coordenação de aperfeiçoamento de pessoal de nível superior brasil (capes) finance code 001. references [1] p. b. fusco. estruturas de concreto: solicitações normais, rio de janeiro: guanabara dois, 1981. [2] american concrete institute aci 318-14. building code requirements for structural concrete (farmington hills: aci), 2014. [3] associação brasileira de normas técnicas. abnt nbr 6118: 2014: projeto de estruturas de concreto procedimento, rio de janeiro: abnt, 2014. [4] a. h. mattock, l. b. kriz, e. hognestad. rectangular concrete stress distribution in ultimate strength design. journal proceedings, 1961. [5] f. n. mendes. bloco retangular versus parábola-retângulo: comparação na flexão composta anais do 4◦ simpósio epusp sobre estruturas de concreto, são paulo: epusp, 2006. [6] j. k. wight. reinforced concrete: mechanics and design, hoboken: pearson, 7th edition, 2016. [7] fib. fédération internationale du béton. model code for concrete structures 2010, lausanne: fib, 2013. https://doi.org/10.1002/9783433604090. [8] european committee for standardization 2004 en 1992-1-1. eurocode2: design of concrete structures part 1-1: general rules and rules for buildings, brussels: cen, 2004. 430 https://doi.org/10.1002/9783433604090 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.35.0054 acta polytechnica ctu proceedings 35:54–60, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague comprehensive verification of the behavior of the continuous welded rail on the bridge jiří vendela, ∗, otto plášeka, jan valehracha, tomáš říhaa, václav machb a brno university of technology, faculty of civil engineering, veveří 331/95, 602 00 brno, czech republic b independent researcher, bílovice, czech republic ∗ corresponding author: xsvendelj@vutbr.cz abstract. the aim of this paper is to verify the real behavior of the continuous welded rail on the bridge with a longer expansion length than allowed by czech national regulations. special attention is focused on the rate of interaction of the continuous welded rail and the bridge, which decisively affects the stress state in the continuous welded rail when the temperature changes. in order to improve the result and increase the accuracy of observation, two measuring methods (geodetic and strain gauge) were chosen while recording the temperature conditions of individual components of the system were simultaneously recorded. keywords: continuous welded rail, longitudinal resistance, thermal expansivity. 1. introduction when designing new bridge structures with a continuous welded rail, it is very appropriate to avoid a solution with expansion devices (or rail joints) in the track for several reasons. this is primarily a benefit of passenger comfort in the case of exclusion crossing of these problem areas, a significant reduction of dynamic effects in transition zone and reduction of maintenance costs. especially when replacing the permanent way on current bridge structures and in adjacent sections of a track is also the need to take into account the financial benefit of establishing a continuous welded rail in comparision with the variant with expansion devices whose acquisition costs are severalfold higher than the costs of a solution with an uninterrupted running edge of the rails. on the other hand the elimination of expansion devices in the bridge zone involve increasing of stress in terms of the combined response of the structure and track dealing with the issue in the design. from the designer’s point of view, however, the current standard provisions may appear as insufficient, especially with regard to the variability of the permanent way system and bridge structures and it happens that certainty is taken into account compromise precautions in the calculations. in aspect of infrastructure serviceability of the continuous welded rail is considered the most important the spacial track stability. its uses on bridges has drawback, therefore it is necessary to take into account especially the span of bridge structures and loads that affect the interacting system. this work solves especially the first part of the problem, which is continuous welded rail and its role in the interaction system with a bridge. it focused particularly on the range of its dependence on a bridge structure movements and phenomena that occur when the state of the whole system changes. the design of the bridge structure and its characteristics deals only marginally, but tries to describe its behavior in detail, especially when temperature conditions changes. a number of research projects focus on determining the combined response of the bridge structure and continuous welded rail from individual types of additional stresses in the rails, which are: • stress from thermal changes in rails, • stress from thermal changes in bridge structure, • stress from traction and braking forces, • stress from classified vertical traffic loads. 2. methodology 2.1. basic assumptions for a theoretical model of a cwr and bridge when calculating the force effects in a track that is only affected by a load from a temperature changes, is based on knowledge of hooke’s law and physical law of thermal expansion material, see equation 1. nx = ea( du dx −α∆t) (1) the system of horizontal springs simplifies the problem of the rigidity of the connection between bridge and rails (see figure 1). the system is divided into individual beams, for which the basic differential equation were defined. the beginnings of the coordinate systems are always on the left sides of the beams, except beam no. 1, which starts on the right side to maintain continuity computational model. differential equations for beams marked with numbers 1, 2, 3, 4 in figure 1 can be written as: 54 https://doi.org/10.14311/app.2022.35.0054 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 35/2022 example of an article with a long title figure 1. a simple girder with a fixed bearing on one side of the bridge. −eiaiu ′′ i + kiui = 0, i = 1, 3, −e2a2u ′′ 2 + k2(u2 −u4) = 0, −e4a4u ′′ 4 + k4(u4 −u2) = 0. (2) and forces in beams are calculated according to equation 1 as: nx = eiai(u ′ i −α∆t). (3) if we want to solve a system of differential equations, it is necessary to determine the boundary conditions. these are compiled based on knowledge of the behavior of the real system, where with growing distance from the ends of the bridge, the influence of the continuous welded rail by the bridge structure decreases. that is, the displacements at a sufficient distance from the bridge are zero and also the displacements and the stresses at the junction of the two beams are the same. then the general solution of the system of differential equations 2 can be written as: ui(x) = bieλix + cie−λix, i = 1, 3, u2(x) = b2eλ2x + c2e−λ2x + α4∆t4x. (4) the exponential term λ takes into account the nonlinear dependence of the longitudinal resistance on the mutual displacement of the track owing to the bridge structure and is calculated according to the following relation: λ2i = ki u0eiai (5) a graphical representation of this nonlinear phenomenon can be seen in the figure 2. uic code 774-3[1] specifies the value of the displacement u0 at the interface of the longitudinal linear and shear resistance and provides at least a basic overview of when approximately occurs shear behavior between the bridge and the continuous welded rail (cwr) when loaded by external forces: • ballasted-deck bridges: displacement at the interface of longitudinal linear and shear resistance: u0 = 0, 5 mm for the resistance of the rail to sliding relative to the sleeper (frozen ballast), u0 = 2 mm for the resistance of the sleeper in the ballast. current values of the resistance in the plastic zone: k = 12 kn.m−1 resistance of sleeper in ballast (unloaded track), moderate maintenance, k = 20 kn.m−1 resistance of sleeper in ballast (unloaded track), good maintenance, k = 60 kn.m−1 resistance of loaded track or track with frozen ballast. • unballasted bridges: displacement at the interface of longitudinal linear and shear resistance: u0 = 0, 5 mm for the resistance of the rail due to the upper surface of the structure. current values of the resistance in the plastic zone: k = 40 kn.m−1 resistance of the rail due to the upper surface of the structure (unloaded track), k = 60 kn.m−1 resistance of the rail due to the upper surface of the structure (loaded track). in laboratory conditions, a number of research projects have demonstrated the validity of the scope values track resistance on bridge structures. but as other authors prove on experimental measurements directly on bridges, these recommendations may not always be generally valid and each new combination of parameters affecting the interaction cwr and the bridge can be determined by direct measurements that the degree of interaction can deviate from the submitted standards. therefore, even this work tries to set values that will capture the bahavior of a particular interacting track and bridge system by experimental research. 3. experimental measurements 3.1. bridge and cwr parameters bridge due to the fact that these are three independent structures whose central longitudinal joints are covered with a metal plate always welded to only one of them, so that they can move independently, only the monitored construction nk1 with the following parameters that are important to describe the behavior of the whole system will be presented below: • marking: nk1, • description: plate steel girder, • type: continuous with 5 fields, • lenght: 80,3 m, • width: 9,5 m, 55 j. vendel, o. plášek, j. valehrach et al. acta polytechnica ctu proceedings l on gi tu di na l sh ea r f or ce [ n ] displacement of the rail u [m] (3) (2) (4) (1) u0 u0 figure 2. variation of longitudinal shear force with longitudinal track displacenlellt for one track. (1) resistance of the rail in sleeper (loaded track without ballast); (2)resistance of sleeper in ballast (loaded track); (3) resistance of the rail in sleeper (unloaded track without ballast); (4) resistance of sleeper in bal1ast (unloaded track). • span: 79,6 m, • lenght: 80,3 m, • number of tracks: 2. continuous welded rail there are three different tracks on the bridge, of which two are double-track and one single track (see figure 3). cwr is set up in all tracks on the bridge. the length of the supporting structure was adjusted to meet the requirements of regulation [2] for uic 60 rails. however, it is located on the nk1 structure track with rails 49 e1 without any special modifications to the railway superstructure. permanent way in this track is in the following composition in the entire length of the bridge and in the adjoining sections: • rails: 49 e1, • concrete sleepers: b 91s/2, • elastic rail fasteners: w 14, • sleepers spacing: 600 mm, • deep trackbed. 3.2. phases of measurement the whole experiment was divided into two phases, which took place independently of each other: • phase i: geodetic measurements of the position of the bridge structure and continuous welded rail based on geodetic monitoring of bridge structure displacement and continuous welded rail depending on temperature changes. • phase ii: tensometric measurements of the stress state of the continuous welded rail based on monitoring the stress state of continuous welded rail by strain gauges depending on temperature changes. 3.3. phase i: geodetic measurements of the position of the bridge structure and continuous welded rail the method of geodetic monitoring with simultaneous measurement of air, rail and bridge temperature was chosen to monitoring the movement of the bridge structure and the continuous welded rail. the following was monitored on the bridge structure in phase i: • displacement of rail strings depending on rail temperature, • displacement of the bridge load-bearing structure nk1 depending on its temperature. the first step was to select the location of the measuring points, marked on both rail strings a total of 20 measuring profiles (20 points on the left and right rail string, numbered 1 to 20 against the direction of stationing). for monitoring displacements of the bridge structure 8 points were fitted (marked m1 to m8 points m1 and m8 were located on abutments, other points on the bridge deck in bridge piers areas). 56 vol. 35/2022 example of an article with a long title figure 3. view of the track on the bridge in the direction from fixed bearing. for monitoring the expansion behavior of the bridge structure and the permanent way depending on the temperature changes were marked observed points in selected cross-section profiles of rail strings whose positional changes in the direction of the track axis were measured in stages, together with the monitoring of other factors (especially the temperature of the rails and the bridge structure). their stabilization was accomplished a mild hole φ 1 mm punched out on the outer unmoved edge of the rail heads. these points were numbered and marked in color on the rail web. applies to both rail strings. 3.4. phase ii: tensometric measurements of the stress state of the continuous welded rail the aim of phase ii was to supply the geodetic monitoring of bridge and track displacements with knowledge in terms of the stress state of the continuous welded rail. during seven continuous measuring cycles, when one cycle always lasted for 3-4 days, data were collected from the installed resistance strain gauges and temperature sensors. recording period of continuously measured quantities was 1 second. in phase ii, data were collected from installed resistance strain gauges and temperature sensors about: • relative deformation of rail strings depending on temperature change, • temperature of the bridge structure and continuous welded rail. from the data obtained from the individual phases, it was necessary to approximate the set of measured data by the given equation so that the approximation function was best matched to the measured data. using force of the matlab program, the parametric space was searched, discretized and seeks the optimum of the function on it. it was necessary to determine the size of the error that occurred during approximation. this was done by calculating the residue, whose sum was minimized to estimate the parameter of the regression function. the residue calculation was used: ei = yi − ŷi (6) and to estimate the regression function parameter by minimizing the residual sum using the following relationship: s2e = n∑ i=1 (yi − ŷi)2 (7) where yi are particular displacement values determined by measurement, ŷi are the displacement values obtained when searching for the minimum of the function on the parametric space of an unknown quantity. 4. results and disscusion 4.1. determination of the equivalent coefficient of thermal expansion of the bridge construction as no data were known on the stiffness of the substructure, friction in moving bearings, nor other possible effects that could also affect the longitudinal expansion bridge structure, therefore the whole phenomenon was expressed by an equivalent factor thermal expansion αm (similar to [3]), which will be determined on the basis of data recorded in phase i. all data (longitudinal displacements) obtained by experimental measurements during nine stages (e0 e8) were used to approximate the finding coefficient solution αm. the data were interpolated with a previously known approximation function: um = lmαm∆tm (8) the graph of this function is a straight line. to solve this linear approximation mathematicalstatistical least squares method (lsm) was chosen. 57 j. vendel, o. plášek, j. valehrach et al. acta polytechnica ctu proceedings figure 4. residual sums of control functions with variable αim on the searched parametric space. figure 5. the resulting value of αm as a minimum of control functions. in general, it was a matter of approximating the set of measured data by the given equation so that the function matched the measured data as best as possible (see figure 4). the resulting value of αm is shown on the figure 5 as the average of the mean square errors of the functions of the individual stages over all iterated αim. the resulting coefficient of thermal longitudinal expansion of the bridge structure, based on geodetic measurements and mathematical linear approximations using least square method, was determined by the value: αm = 9, 66 · 10−6 k−1 . looking into recent experimental research (for example experimental measurements on the znojmo viaduct or on the bridge in kolín) it can be stated that for these three steel structures with a rail bed the bridge structures are comparable and very similar in terms of the determined equivalent coefficient of thermal expansion of the bridge structure (znojmo – αm = 9, 7 · 10−6 k−1, kolín – αm = 8, 5 a 8, 9 · 10−6 k−1). figure 6. residual sums of control functions with variable kiip on the searched parametric space. figure 7. the resulting kip value from phase i as a minimum of control functions. 4.2. determination of the longitudinal resistance of the track on the bridge based on the principles described in chapter 2.1, approximation was performed on experimental data. in this case, it is a general linear regression using general polynomials defined by differential equations 4. for solutions in this case, too, the lsm was chosen, which helped with sufficient accuracy determine the finding values. in accordance with the requirements [4] the influence of the longitudinal shear resistance kip according to the figure 2 was taken into account. the approximation was further carried out by the above equations 6 and 7, it means that after the search of the parametric space, the resulting optimized function reached the lowest residual sum. the area where the minimum mean square error of the control equations was expected is shown by a black bar in figure 6. the resulting value of longitudinal shear resistance of the track kip is shown in figure 7 as the average of the mean square errors of the functions of individual stages over all iterated kiip. the resulting longitudinal resistance of the track 58 vol. 35/2022 example of an article with a long title figure 8. residual sums of control functions with variable kiip across all recordsets. on the bridge without loading by rail traffic, based on geodetic measurements of the position of the rail strings and mathematical general linear approximation using least square method, was determined by the value: ki = 4, 5 · 103 n.m−1 . 4.3. longitudinal resistance cwr on the bridge according to tensometric measurements this chapter will verify the conclusions from chapter 4.2, where the value of the longitudinal resistance cwr was determined on the bridge. the following evaluation was performed on the based of strain gauges measurement of relative deformation of rail string and measurement temperature of the bridge construction and track. a total of 1, 442, 033 values were recorded in 7 measurement cycles from each strain gauge and the same number of temperature sensors. the number of values corresponds to total measurement length in seconds. for a purpose the same method for the evaluation was applied to as a reliable comparison of the results data from geodetic measurements, it means by minimizing the residual sum of deviations measured and approximated by lsm with the validity of equations 6 and 7. the course of residual sums of control functions with the variable kiip across all recordsets are seen in the figure 8. the approximation functions are the equations of the individual axial forces. parametric space in which the solution of the longitudinal shear resistance value was searched, was limited as follows: • n〈1; 1442025〉 is the number of records, • kiip〈1; 10000〉 is the range of the vector of the searched longitudinal shear resistance of the track, • i〈1; 14〉 is the number of monitored strain gauges. figure 9. the resulting value of kiip from phase ii as a minimum of control functions. in this space, based on tensometric measurements of the relative deformation of the left rail string and temperature measurements of the bridge structure and the track, it was found out that there was a good agreement with the results obtained by geodetic measurements. the right rail string is assumed to behave symmetrically in terms of tension. therefore with a degree of simplification, based on the assumption of symmetry, is the resulting value double of the longitudinal shear resistance found out on the left rail string. by tensometric measurement of the left rail string and mathematical general linear approximation using lsm, was found out the value of longitudinal shear resistance of the track on the bridge without load by rail traffic of the size: ki = 4, 6 · 103 n.m−1 . 5. conclusions it is apparent that the resulting values from both phases are in area of very low values of longitudinal shear resistance of the track. in other words, the tension in the rails due to thermal expansion of bridge structure is much less affected than might be expected. this is evident by the fact that the track on the bridge, with an expansion length of more than 20 m longer than permitted by national regulation [2] in the relevant part xii, does not show, more than ten years of operation, anomalies in the behavior of continuous welded rail and there are no defects on permanent way of a greater extent than anywhere else on the line. acknowledgements project: interakce koleje a mostů s velkými dilatačními délkami; hs12257021210001 in cooperation sždc, s.o. and brno university in 2013 – 2015. specific research: sledování bezstykové koleje na mostě; fast-j-14-2528 in 2016. 59 j. vendel, o. plášek, j. valehrach et al. acta polytechnica ctu proceedings references [1] uic code 774-3r. track/bridge interaction: recommendations for calculations. 2nd edition. paris – france: international union of railways, 2001. 70 p., isbn 2-7461-0257-9. [2] sždc s3. železniční svršek – regulation. effective from 1. 10. 2008 as amended by dmendment no. 4 (effective from 1. 3. 2021). praha: sždc, s.o., 423 p., 2008. [3] l. frýba. dynamika železničních mostů. praha: academia, 1992. 328 s., isbn 80-200-0262-6. [4] čsn en 1991-2 (73 6203). eurocode 1: actions on structures – part 2: traffic loads on bridges. prague: czech committee for standardization, 152 p., 2005. 60 acta polytechnica ctu proceedings 35:54–60, 2022 1 introduction 2 methodology 2.1 basic assumptions for a theoretical model of a cwr and bridge 3 experimental measurements 3.1 bridge and cwr parameters 3.2 phases of measurement 3.3 phase i: geodetic measurements of the position of the bridge structure and continuous welded rail 3.4 phase ii: tensometric measurements of the stress state of the continuous welded rail 4 results and disscusion 4.1 determination of the equivalent coefficient of thermal expansion of the bridge construction 4.2 determination of the longitudinal resistance of the track on the bridge 4.3 longitudinal resistance cwr on the bridge according to tensometric measurements 5 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0140 acta polytechnica ctu proceedings 33:140–145, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague circular concrete: validating new technologies in the lab and on-site bram doomsa, ∗, jeroen vrijdersb a belgian building research institute, division of geotechnics, structures and concrete, rue du lombard 42, 1000 brussels, belgium b belgian building research institute, division of intelligent installations and sustainable solutions, rue du lombard 42, 1000 brussels, belgium ∗ corresponding author: bram.dooms@bbri.be abstract. the growing importance of a circular economy also holds possibilities for concrete. some of the principles of circularity, like the re-use of "waste" as secondary raw material and the optimisation of the design with regards to lifespan, environmental impact and re-use, can be applied. several technological innovations in this domain are in an advanced stage of development. however, the transition from research to full commercial applications is held back, mostly due to the absence of documented experience and the lack of a technical framework (prescriptions). in order to accelerate the adoption of new technologies, the belgian building research institute (bbri) started the project "circular.concrete". in this paper, the results are presented of validation tests regarding 1) smart crushing technology, 2) the carbonation of recycled aggregates, 3) concrete using only recycled aggregates and 4) the use of alternative binders and supplementary cementitious materials. keywords: carbonatation, concrete recycling, pilot projects, recycled concrete aggregates (rca), supplementary cementitious materials. 1. introduction substantiated by the scientific consensus of global warming and climate change and propelled by the increased public awareness, environmental clauses are frequently added to the specifications for buildings for major clients. concrete has a dubious reputation in this matter. this is mostly due to the enormous volumes of cement that are produced worldwide on a yearly basis (∼ 4.5 billion tons, 2019), the co2emissions during cement production (∼ 800 kg of co2 per tonne of cement for cem i), the extraction, processing and transportation of the raw materials and the important waste streams that are produced when concrete structures are demolished [1, 2]. the growing importance of a circular economy also holds opportunities for concrete. some of its principles like the re-use of "waste" as secondary raw material and the optimisation of the design with regards to lifespan, environmental impact and re-use, can be applied. several technological innovations in this domain are in an advanced stage of development. however, the transition from research to real applications is held back, mostly due to the absence of documented experience and of a technical framework, i.e. technical prescriptions, standards etc. in order to accelerate the adoption of new technologies, the belgian building research institute (bbri) started the project circular.concrete. the aim of this project is to translate research results into practice. 2. validating innovative technologies in a first stage of the circular.concrete project, different technological innovations were identified and their trl (technology readiness level) were estimated. the results have been summarized in a stateof-the-art report, which is available on the project website (https://www.circular-concrete.be/). in a second stage a selection was made of technological innovations for further validation on laboratoryscale, based on their trl, their scientific potential and their relevance for the belgian market: 1) smart crushing technology, 2) the carbonation of recycled aggregates, 3) concrete using only recycled aggregates and 4) the use of alternative binders and supplementary cementitious materials. it should be noted that the scale of this validation was very limited and should be considered as exploratory. 2.1. technology 1. smart crushing using the smart crushing technology, concrete construction waste can supposedly be crushed and processed into its original components, i.e. the original fine and coarse aggregates and the binder, of which a part is non-hydrated (and still reactive) [3]. since the obtained recycled aggregates don’t have any mortar adhered, they can be used in new concrete as natural aggregates. this technology was tested on four different belgian concrete waste samples: 140 https://doi.org/10.14311/app.2022.33.0140 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en https://www.circular-concrete.be/ vol. 33/2022 circular concrete: new technologies figure 1. sample 4: low quality concrete waste fraction 20-56 mm. water absorption [%] particle density [kg/m3] micro-deval [−] sample 1 3.8 2440 23 sample 2 6.1 2370 28 sample 3 4.5 2400 23 sample 4 4.3 2420 21 limestone aggregate 0.5 − 1.0 2660 − 2800 15 table 1. characterization of the treated aggregate samples. 1. concrete waste fraction 2-20 cm, originating from railway sleepers, with strength class c50/60 and no reinforcements or contaminants. 2. concrete waste fraction 2-20 cm, originating from the precast industry, with strength class c45/55 and no reinforcements or contaminants. 3. low quality concrete waste fraction 20-56 mm, composed of concrete (70 wt%), natural stone (25 wt%), and masonry and contaminants (5 wt%). 4. low quality concrete waste fraction 20-56 mm, of which individual aggregates can be composed of concrete (80 wt%), fibre reinforced concrete (10 wt%), or asphalt (10 wt%) (figure 1) 500 kg of each sample was sent to a smart crushing facility in the netherlands to be processed. after processing, about 300 kg of material from each sample was returned. 2.1.1. aggregate testing each treated sample was assessed visually and the grain size distribution was determined. all treated samples presented a very similar distribution curve (0/22). the water absorption and particle density (on the fraction +4 mm, in accordance with en 1097-6) and resistance to wear (on the fraction +10 mm, in accordance with en 1097-1) was determined for the treated samples. the results are presented in table 1. the values for the treated samples are very different from the typical values for crushed natural limestone aggregates. these differences can be explained by the presence of mortar on the treated samples, which hasn’t been removed by the smart crushing treatment. belgian concrete is generally made with crushed limestone as coarse material, whereas dutch concrete generally contains natural (round) gravel. possibly this technology is more effective for concrete composed of rounded natural gravel. 2.1.2. fines testing from each treated sample, the fines (< 0.125 mm) were extracted and tested, to determine the potential presence of non-hydrated cement. x-ray diffraction analysis (xrd) of the fines demonstrated that this fraction is dominantly composed of quartz representing aliquots of the silt and sand aggregate fraction of the primary concrete. mortar bars were made (in accordance with en 196-1) using standard sand and cem i 52.5 n, and the fines as cement replacement (40 %). the fresh properties and the mechanical properties of the mortar bars after 7, 28 and 91 days were compared to those of reference mortar bars without cement replacement. the results of the mechanical tests (executed in accordance with en 196-1) are presented in table 2. the results indicate that the fines do not contribute to the strength development of the mortar. possibly the evaluated fraction should be even finer (< 0.063 mm) to exclude as much as possible the presence of silt and sand. 2.1.3. concrete testing using only the treated samples as received from the smart-crushing facility, concrete was prepared with each sample by adding 300 kg/m3 cem iii/a 42.5 and water in a w/c ratio of 0.55. due to an error in 141 bram dooms, jeroen vrijders acta polytechnica ctu proceedings rf,7d rc,7d rf,28d rc,28d rf,91d rc,91d [mpa] [mpa] [mpa] [mpa] [mpa] [mpa] sample 1 5.1 27.1 6.2 31.8 6.9 33.8 sample 2 5.6 27.4 6.4 33.7 6.3 34.7 sample 3 5.6 26.5 5.7 31.3 7.0 33.4 sample 4 5.3 26.9 5.9 32.3 6.2 33.2 reference mix 10.0 63.9 9.4 71.8 10.0 79.3 table 2. mechanical properties of the mortar mixtures with treated fines. rc,28d rc,91d carbonation depth 56 d [mpa] [mpa] [mm] sample 1 21.9 28.8 13.5 sample 2 22.2 23.5 15.5 sample 4 20.3 24.9 17.3 reference mix 30.5 45.8 11.3 table 3. compressive strength and carbonation resistance of the concrete mixtures with treated aggregates. the mix design, a deviant concrete composition was obtained with sample 3, which lead to the rejection of this concrete mixture. without adding superplasticizer but taking into account the water absorption of the treated samples, a very good workability was obtained of the concrete mixtures. the compressive strength after 28 and 91 days (determined in accordance with en 12390-3), and the resistance to carbonation (determined in accordance with en 13295) were determined for the mixtures and compared with a reference concrete mixture, using only crushed natural limestone aggregates. the composed grain size distributions of the different concrete mixtures differ slightly. the results are presented in table 3. a clear influence can be observed of de total replacement of natural aggregates with the treated samples. after 28 days a strength reduction of about 30 % can be observed, after 91 days of about 40-50 %. the concrete with treated sample 3 shows deviant behaviour. the carbonation depth of the concrete after 56 days of exposure to 1.0 % co2 was about 20-50 % higher for the treated samples in comparison to the reference mix. 2.2. technology 2. carbonation of recycled aggregates with the carbonation of recycled aggregates, their physical and mechanical properties can be enhanced and thus their applicability in concrete improved. at the same time co2 can be ’captured’ inside the aggregates. recycled aggregates of different origins (concrete ’c’, high quality concrete ’c a+’, mixed ’m’) and of different sizes (0/6.3, 6/22, 8/20, 0/40 and 10/40) were treated during 5 days in an autoclave at an increased pressure (2 bar) and a very high co2 concentration (100 %). before and after the treatment, the water absorption and particle density (in accordance with en 1097-6) and resistance to wear (in accordance with en 1097-1) were determined for the aggregates. the results are presented in table 4. the co2-treatment clearly affected the water absorption and the particle density of the aggregates. as only the mortar fraction, adhered to the recycled aggregates, is influenced by the carbonation, the aggregate properties are affected proportionally in relation to the mortar content of the aggregates. 2.3. technology 3. maximum aggregate replacement by replacing the natural aggregates in concrete by recycled or artificial aggregates, the need for primary raw aggregates can be reduced. the influence of the presence of recycled aggregates (ra) and artificial aggregates on the properties of fresh and hardened concrete was evaluated by replacing a maximum quantity of natural aggregates in concrete mixtures (% vol). this maximum quantity was determined by combining the recycled and artificial aggregates as received with natural aggregates, while fitting the combined grain size distribution of the concrete to a reference grain size distribution. the following aggregate combinations were tested: 1. 10 % 0/1 natural sand + 90 % ra type 1 (0/8 + 8/22) 2. 100 % ra type 2 (0/8 + 6/20) 3. 10 % 0/1 natural sand + 90 % ra type 3 (0/8 + 6/20) 4. 10 % 0/1 natural sand + 26 % 0/4 treated polluted sand + 64 % crushed limestone (4/20) 5. 20 % 0/1 natural sand + 20 % 0/4 copper slag + 60 % crushed limestone (4/20) 142 vol. 33/2022 circular concrete: new technologies water absorption particle density micro-deval [%] [mg/m3] [−] before after before after before after 0/40 c 6.8 5.4 2.20 2.29 25 27 10/40 m 7.6 7.4 2.04 2.08 42 47 0/6.3 c 6.3 5.9 2.29 2.33 / / 8/20 c a+ 4.3 3.7 2.42 2.44 16 13 8/20 c 4.4 3.5 2.37 2.43 24 16 6/22 c 4.3 3.8 2.34 2.37 11 12 table 4. characterization of the treated aggregates. type 1 concrete type 2 concrete rc,28d rc,91d carbonation depth 56 d rc,28d rc,91d carbonation depth 56 d [mpa] [mpa] [mm] [mpa] [mpa] [mm] combination 1 30.3 33.8 14.8 40.0 36.0 9.8 combination 2 26.6 25.9 17.5 32.8 32.3 14.0 combination 3 29.7 31.4 13.8 37.2 37.8 10.5 combination 4 40.9 53.1 11.0 51.3 58.0 6.3 combination 5 32.7 44.4 11.3 43.2 61.8 8.8 combination 6 29.9 41.4 12.0 57.6 61.7 7.5 reference mix 30.5 45.8 11.3 55.9 62.8 7.3 table 5. compressive strength and carbonation resistance of the concrete mixtures with maximum aggregate replacement. 6. 10 % 0/1 natural sand + 50 % inox steel slag (0/2 + 2/10) + 40 % crushed limestone (10/20) two types of concrete were made with those aggregate combinations: type 1 with 300 kg/m3 cem iii/a 42.5 and a w/c ratio of 0.55, and type 2 with 340 kg/m3 cem iii/a 42.5 and a w/c ratio of 0.45. superplasticizer was added to the fresh concrete to obtain an initial consistence class s3 s4 (in accordance with en 206). the results were compared to a reference concrete mixture using only crushed natural limestone aggregates. it is important to note that, while all concrete mixtures were designed using the same reference grain size distribution for the aggregates, differences exist between the actual grain size distributions. in addition, all the aggregates were applied in oven-dried state, and extra water was added to take into account their total water absorption. considering the differences in the kinetics of the water absorption of the aggregates, the effective w/c ratio of the mixtures may be higher than the targeted values. this is particularly the case with combinations 1 to 3, given the high replacement rates and high water absorption of recycled aggregates. this should be taken into account when considering the results presented in this article. the compressive strength after 28 and 91 days (determined in accordance with en 12390-3), and the resistance to carbonation (determined in accordance with en 13295) were determined of the concrete mixtures. the results are presented in table 5. for type 1 concrete, the influence of the maximum aggregate replacement on the compressive strength after 28 days is very low. only the concrete with combination 4 seems to perform significantly better. after 91 days however, a clear influence on the compressive strength can be observed in the case of maximum aggregate replacement by ra (combinations 1 to 3), with reductions of about 25-45 %. the carbonation resistance of these ra concrete mixtures is also clearly lower than that of the reference mix. for type 2 concrete, the maximum replacement of aggregates by ra (combinations 1 to 3) leads to a reduction of the compressive strength after 28 days of about 30 − 40 % and after 91 days of 40-50 %. the carbonation resistance of these ra concrete mixtures is also significantly lower than that of the reference mix. the concrete with combinations 4 to 6 behaves similar to the reference concrete mix. the freeze-thaw resistance was determined in accordance with cen/ts 12390-9 of the hardened concrete mixtures (only type 2 concrete). the results are presented in table 6. overall, the influence of the maximum aggregate replacement on the freezethaw resistance does not seem to be important. only the concrete with combination 2 (absence of natural sand) clearly performs worse than the other concrete mixtures. the air content of this concrete mixture (1.6 %) was lower than that of the other mixtures (2.0 − 4.8 %), which probably contributed to this re143 bram dooms, jeroen vrijders acta polytechnica ctu proceedings mean cumulative scaled material after 7 cycles 14 cycles 28 cycles 42 cycles 56 cycles [kg/m2] [kg/m2] [kg/m2] [kg/m2] [kg/m2] combination 1 0.74 1.02 1.30 1.58 1.82 combination 2 1.30 2.08 3.04 4.38 5.54 combination 3 0.28 0.68 1.28 1.76 2.28 combination 4 0.14 0.50 1.06 1.50 1.98 combination 5 0.66 0.96 1.60 2.42 3.00 combination 6 0.50 0.80 1.18 1.70 2.18 reference mix 0.58 0.94 1.42 1.90 2.46 table 6. freeze-thaw resistance of the concrete mixtures with maximum aggregate replacement. sult. the resistance to chloride migration was determined in accordance with nt build 492 of the hardened concrete mixtures (only type 2 concrete). the results are presented in table 7. the maximum replacement of aggregates by ra (combinations 1 to 3) seems to lead to a reduction of the resistance to chloride migration. again, the concrete mixture with combination 2 (absence of natural sand) clearly performs worse than the other ra concrete mixtures. the mixtures with slag aggregates (combinations 5 and 6) seem to perform slightly better than the reference mix. non steady state chloride migration coefficient dnssm [kg/m2] combination 1 6.43 combination 2 10.17 combination 3 8.46 combination 4 6.57 combination 5 4.90 combination 6 3.98 reference mix 5.98 table 7. resistance to chloride migration of the concrete mixtures with maximum aggregate replacement. 2.4. technology 4. alternative binders and supplementary cementitious materials by replacing the traditional portland cement in concrete completely by alternative binders or partially by supplementary cementitious materials, the co2 emissions related to concrete production can be reduced substantially. four alternative binders (’c’, replacing 100 % of traditional cement) and five supplementary cementitious materials (’f’, replacing 40 % of the traditional cement) were selected. c1 and c2 are sulphoaluminate cements, currently available on the market. c3 is a cem iii/a 42.5 cement and c4 a cem iii/b 32.5 cement. f1 is a copper slag, f2 and f3 are inox slags and f4 is a granulated blastfurnace slag. f5 consists of calcinated clay. with cem i 52.5 n as reference cement type and using standard sand, mortar mixtures conforming to en 196-1 were made with the alternative binders and supplementary cementitious materials. the setting times of the mortar mixtures were evaluated conform en 480-2. the results are presented in table 8. depending on the binder type, the alternative binders present a faster (c1 and c2) or a slower (c3 and c4) setting than the reference portland cement. the supplementary cementitious materials all have slower setting times than the reference cement. initial setting time final setting time [minutes] [minutes] reference mix 214 278 c2 103 145 c3 297 403 c4 358 584 f1 293 398 f2 321 418 f3 312 420 f4 267 382 f5 279 388 table 8. resistance to chloride migration of the concrete mixtures with maximum aggregate replacement. the mechanical properties of the mortar mixtures with the alternative binders and supplementary cementitious materials were evaluated in accordance with en 196-1. the results are presented in table 9. all the alternative binders and supplementary cementitious materials lead to lowered mechanical properties of the resulting mortar. 3. conclusions the circular.concrete project aims to identify interesting innovative technologies in the domain of circular economy and enable or broaden their practical application. four different innovative technologies were 144 vol. 33/2022 circular concrete: new technologies rf,7d rc,7d rf,28d rc,28d rf,91d rc,91d [mpa] [mpa] [mpa] [mpa] [mpa] [mpa] reference 10.0 63.9 9.4 71.8 10.1 79.3 c1 5.7 39.5 6.3 43.4 7.0 45.2 c2 8.6 45.3 7.4 50.1 6.5 50.1 c3 6.2 30.4 7.5 42.8 9.6 56.6 c4 5.6 23.4 7.2 34.1 7.3 40.1 f1 6.3 33.0 5.9 35.0 5.6 37.1 f2 5.9 27.2 6.4 35.3 7.0 37.1 f3 4.9 21.4 5.3 26.7 6.0 31.8 f4 5.8 34.4 8.1 49.5 6.5 54.7 f5 6.0 29.0 6.7 40.0 7.7 47.3 table 9. mechanical properties of the mortar mixtures with alternative binders and supplementary cementitious materials. selected and validated in laboratory-scale tests in order to increase their trl. it has to be noted that the scale of this validation was very limited and should be considered as exploratory. the smart crushing technology, which supposedly crushes the concrete into its original components, doesn’t seem very effective for belgian concrete, made with crushed limestone aggregates. the crushed secondary aggregates still contained a lot of adhered mortar. possibly this technology is more effective with concrete with rounded natural gravel. the returned fines (< 0.125 mm) did not show hydraulicity as they are dominantly composed of quartz. the carbonation of recycled aggregates has a clear positive, albeit not that important, influence on their physical and mechanical properties. replacement rates of natural aggregates in concrete of 90-100 % vol by recycled aggregates logically lead to reduced mechanical properties of the concrete (this effect becomes more important for higher concrete strength classes). the durability properties of the concrete are also reduced, although this cannot be concluded for the freeze-thaw resistance (the air content of the concrete seems to be the determining factor for this property). the partial replacement by artificial aggregates did not strongly impact the concrete properties. all the evaluated alternative binders (replacing 100 % of the traditional cement) and supplementary cementitious materials (replacing 40 % of the traditional cement) lead to lowered mechanical properties of the resulting mortar. this research demonstrates the importance of validating innovative technologies by independent laboratories before their application in concrete projects. one the one hand, claims can be validated, on the other hand, extra experience is being built to create the best application conditions in real scale projects. acknowledgements the authors would like to thank the vlaams agentschap innoveren en ondernemen (vlaio), sim (strategic initiative materials) and a great number of industrial partners for their financial support of the circular.concrete project. (www.circular-concrete.be). references [1] cembureau. environmental product declaration (epd) according to en 15804 and iso 14025 portland cement (cem i) produced in europe, 2020. https://cembureau.eu/policy-focus/sustainable-c onstruction/sustainability-standards/. [2] j. lehne, f. preston. making concrete change innovation in low-carbon cement and concrete, chatham house report, 2018. https://www.chathamhouse.org/sites/default/fil es/publications/2018-06-13-making-concrete-cha nge-cement-lehne-preston-final.pdf. [3] smartcrusher bv. from design of the next generation of economical and ecological equipment to delivery and maintenance, 2020. https://www.slimbreker.nl/. 145 www.circular-concrete.be https://cembureau.eu/policy-focus/sustainable-construction/sustainability-standards/ https://www.chathamhouse.org/sites/default/files/publications/2018-06-13-making-concrete-change-cement-lehne-preston-final.pdf https://www.slimbreker.nl/ https://doi.org/10.14311/app.2022.33.0585 acta polytechnica ctu proceedings 33:585–590, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague environmental assessment of non-metallic reinforcement for concrete structures as an alternative to steel reinforcement nadine stoiber∗, mathias hammerl, benjamin kromoser research group for resource-efficient structural engineering, university of natural resources and life sciences vienna, peter-jordan-straße 82, 1190 vienna, austria ∗ corresponding author: nadine.stoiber@boku.ac.at abstract. the concrete industry accounts for a significant amount of co2 emissions worldwide. one approach to counter this issue includes material reduction of structural components via the use of nonmetallic reinforcement, such as carbon, glass and basalt fibre reinforced polymers. on the one hand, non-metallic reinforcement. however, as its environmental impact has not been sufficiently investigated yet, a life cycle assessment of the production phase is presented within this paper. in a first step, the environmental impact of the sole various reinforcement components and types is compared to each other per mass, per tensile or rather yield strength as well as density unit, at which an environmental disadvantage of especially carbon-fibre reinforced polymers is apparent in most cases. in a further step, a focus is put on applying the environmental data of carbon-fibre reinforced polymers to a pedestrian bridge, which is finally compared to a conventionally reinforced concrete bridge and a steel bridge with similar boundary conditions. the latter results indicate that an adequate application of carbon-fibre reinforcement in structural components has the potential to lead to designs of less environmental impact in comparison to conventionally reinforced pendants. keywords: concrete structures, environmental impact, frp reinforcement, life cycle assessment, non-metallic reinforcement. 1. introduction: ’concrete footprint’ according to [1], in 2018 the construction industry was responsible for approximately 10% of co2 emissions worldwide. on a global level, the report further identifies concrete as most widely used building material type. what is more, it is predicted by [2] that the building material concrete will account for around 12% of worldwide greenhouse gas emissions in 2060. besides emissions regarding the production phase and respective required fuels, concrete is characterised by so-called process emissions due to its intrinsic chemical reduction from calcium carbonate to calcium oxide, which represent around two thirds of total emissions [3]. these numbers outline the urgency as well as the potential of the concrete industry in pushing attempts to decrease its emission-intensity forward. several options exist to counter the above outlined issue, at which three major approaches, according to [4], are briefly listed: reducing the amount of cement via an optimization of the concrete mixture, a reduction of concrete material used via a structural optimization of the design as well as the use of high-strength materials such as ultra-high performance concrete (uhpc) and fibre-reinforced polymers (frp) made out of basalt, glass or carbon fibres. the latter approach is pursued within this paper, at which the environmental performance is subject of investigation. firstly, the method and materials are outlined. secondly, the investigated reinforcement types are compared with each other, solely on material level. here, several matrix as well as fibre types are analysed. subsequently, the application of the environmental data of carbon fibre-reinforced polymers to pedestrian bridges as conducted in [5] is presented. 2. methodology: life cycle assessment according to en iso 14050, life cycle assessment (lca) is described as an assessment of the inputs, outputs and the potential environmental impacts of a product system throughout its life cycle. lca is divided into several phases, including life cycle inventory (lci) analysis, where the inputs and outputs of a product system are determined, as well as life impact assessment analysis (lcia), where the environmental impacts of a product system are evaluated and assessed. several characterisation models exist regarding the latter step of an lca, at which cmlia from the university leiden is mentioned within this context [6]. the underlying standardisation of conducting an lca comprises en iso 14040 (fundamentals) and en iso 14044 (instructions). en iso 14025 outlines the use of life cycle assessment to establish an envi585 https://doi.org/10.14311/app.2022.33.0585 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en j. anderson, m. hammerl, b. kromoser acta polytechnica ctu proceedings figure 1. comparison of gwp values extracted from epds with the european average of [7]. ronmental product declaration (epd). the national standard önorm en 15804 has to be highlighted in this context, which regulates the development of epds for the product category of construction products. life cycle phases of a product system range from product, construction process to the end of life stage. the product stage is compulsory to be covered when developing an epd, at which this stage is divided into three categories: raw material supply (a1), transport (a2) and manufacturing (a3). this type of lca is described as cradle-to-gate. commonly, a cradle-to-gate lca represents the most adequate type, as the inputs and outputs are often characterised by satisfying data completeness and quality. other life cycle stages, such as the construction process stage, are accompanied by comparatively more underlying assumptions. at best, all life cycle stages are covered within an epd. due to limited data availability, the production stage of the considered construction products is assessed within this paper. furthermore, several impact indicators are selected to be evaluated, such as the global warming potential (gwp) in kg co2-eq., the acidification potential (ap) in kg so2-eq. as well as the abiotic deletion potential of fossil fuels (adpf) in mj. commonly, due to public popularity, the gwp is paid the most attention to. in order to establish a wider picture of the environmental impact of the considered products, further impact indicators should be assessed. consulted sources within this paper include epds and life cycle inventory data. details about the respective sources as well as the chosen declared or rather functional units are outlined in the sections below. it is crucial to note that irregularities go hand in hand with the comparison of environmental data from varying sources. for example, the lcia phase of the life cycle assessment can be based on varying characterisation models. the authors aimed at maintaining maximum transparency within the process to enable traceability and reproducibility. 2.1. excursion: data bias to briefly outline the phenomenon of data bias and raise awareness regarding data quality within life cycle assessment, the example of steel reinforcement is consulted. extracted gwp values of steel reinforcement (cradle-to-gate lca, declared unit: 1 kg) as done by [5] are illustrated in figure 1. the first five, grey bars represent extracted gwp values from european epds, at which the last, green bar outlines the european average of steel production in 2019 provided by [7]. for further detailed information on data sources of the epds, the reader is directly referred to [5]. given the gwp values extracted from the epds in figure 1, one would assume to have an average value of around 1.0 kg co2-eq., nevertheless, the european average is higher. when analysing the manufacturing methods of the various epds it becomes apparent that, except for ukraine, all epds are built up on electric arc furnace. nevertheless, [8] shows otherwise, at which not electric arc furnace but basic oxygen steelmaking represents the major manufacturing method of steel products in europe. the average of the given data extracted from the epds is therefore not representative for the global warming potential of steel reinforcement in europe. 3. results on material level: comparison of reinforcement components and types figure 2 and figure 3 show the normalized results of the gwp of various fibre and matrix types. gwp values for carbon fibres are extracted from [9], for glass fibres from [10], for basalt fibres from [11], for styrene butadiene rubber and acrylate from [12], for polyester resin, vinylester resin and epoxy resin from [13] as well as from [14]. impact indicators, like the gwp, were directly extracted from these sources. only the last source, [14], represents lci data, which was extracted and subjected to selfevaluation according to the specifications of önorm en 15804:2020. in all cases, the production stage (cradle-to-gate) was considered. regarding figure 3 it is apparent that carbon fibres are characterised by the highest global warming potential in comparison to glass and basalt fibres. as reason the energy-intense production process up to 3,000řc is mentioned. the differences between the various matrix materials in figure 2 are not that significant, nevertheless, a row in dependence of the gwp is possible. the normalized gwp values of carbon fibre-reinforced polymers (cfrp), glass fibre-reinforced polymers (gfrp) and basalt fibre-reinforced polymers (bfrp) are illustrated in figure 4, at which epoxy resin was chosen as matrix material. the results are presented in a normalized 586 vol. 33/2022 non-metallic reinforcement for concrete structures figure 2. comparison of the gwp of different matrix materials (normalized values). figure 3. comparison of the gwp of different fibre types (normalized values). figure 4. comparison of the gwp of different frps (normalized values). manner to enhance readability. the respective components of the frps were combined according to an analysis of data sheets, at which the respective fibre share ranges from 75 to 90% and the matrix material share ranges from 10 to 25%. the gwp of carbon fibres of 26.4 kg co2-eq. per kg was extracted from [9] and is given as example. comparing figure 3 with figure 4 it can be seen that the trend of carbon being the least favourable material choice in case of environmental impact continues, nevertheless the difference between the varying fibre types becomes smaller. in this context, the study of [15] is mentioned, who conducted a cradle-to-grave lca for conventional steeland textile-reinforced (glass, carbon and basalt) façade elements. their findings indicate that all textile versions show comparatively less environmental impact. the carbon version showed a comparatively higher impact than the glass or basalt solution amongst most of the investigated impact indicators (gwp, ap and eutrophication potential). 3.1. in detail: cfrp reinforcement the environmental impact of cfrp is further investigated based on evaluations done in [5], at which the analysis is extended to impact indicators such as the acidification potential in kg so2-eq. (ap) and the abiotic depletion potential of fossil fuels in mj (adpf). furthermore, the environmental impact of cfrp reinforcement is directly compared to the one of conventional steel reinforcement (european average, compare with figure 1). the results are illustrated in figure 5 to figure 7. the properties of a cfrp rebar product from solidian gmbh are consulted [16]. in figure 5, the results are solely based on masses or rather per declared unit of 1 kg of the product system. this figure shows the environmental advantage of conventional steel reinforcement in comparison to cfrp reinforcement amongst all considered impact indicators. figure 6, which is based on masses as well as on the performance of the reinforcement type, still shows environmental benefits of steel reinforcement. nevertheless, an additional consideration of the performance, more precisely the tensile strength of 2,100 mpa of cfrp reinforcement as well as the yield strength of 550 mpa of steel reinforcement shows results more on favour of cfrp reinforcement than a sole consideration of the masses as illustrated in figure 5. figure 7 shows otherwise, at which an environmental benefit of cfrp reinforcement amongst the impact indicators gwp and adpf is visible. here, the performance (tensile and yield strength) as well as the related densities (7,850 kg/m3 for steel reinforcement, 1,500 kg/m3 for cfrp reinforcement) were considered. figure 7 already shows the potential of cfrp reinforcement as an environmentally friendly alternative to steel reinforcement on the material level. 587 j. anderson, m. hammerl, b. kromoser acta polytechnica ctu proceedings figure 5. comparison of steel and cfrp reinforcement per kg product system (normalized values). figure 6. comparison of steel and cfrp reinforcement per kg product system as well as per mpa tensile respectively yield strength (normalized values). 4. results on system level: comparison of pedestrian bridges the environmental data of cfrp were further applied to pedestrian bridges of varying building types by [5], whose results will be outlined as a subsequent step. the consulted bridges are illustrated in figure 8. the first bridge is a carbon concrete bridge in albstadt ebingen, germany and was built in 2015 (abbreviated with b1-ccb), which is characterised by a thin cfrp reinforced cross section and a galvanized steel railing. the second bridge is a conventional steel reinforced concrete bridge, built in 2009 in vienna, austria (abbreviated with b2-rcb). the last bridge, a mild steel bridge, was built in 1999 in vienna, austria (abbreviated with b3-sb). the masses for the quantity survey were extracted from [17] for b1-ccb and provided by the municipal department 29 in vienna austria in case of b2-rcb and b3-sb. further details can be found in [5]. the functional unit for the subsequent comparison is a pedestrian bridge with a span length of ∼ 15 m, an effective width of ∼ 3 m and an imposed load of ∼ 5 kn/m2. the materials of the superstructure were considered. the results of the environmental assessment of the pedestrian bridge types are illustrated in figure 9. the impact indicators gwp and adpf are clearly in favour of the cfrp reinforced concrete bridge due to figure 7. comparison of steel and cfrp reinforcement per mpa tensile respectively yield strength as well as per density unit (normalized values). significant savings in material masses when comparing b1-ccb with b2-rcb. to get a deeper understanding of the allocation of the environmental impacts, the results are furthermore divided per building material and outlined in figure 10 to figure 12. the consideration of the gwp as well as of the adpf shows that the galvanized steel railing of b1-ccb has a significant share in the whole environmental impact. regarding adpf, the reinforcement is responsible for the majority of the environmental impact. regarding ap in figure 11, it becomes apparent that the cfrp reinforcement is mainly responsible. the environmental data are characterised by high limits of variation, at which [9] gives the consultation of different reaction equations for the combustion of ammonia and hydrocyanic acid during the production process of carbon fibres as reason for these variabilities. this circumstance outlines room for higher data precision regarding the ap in the future. furthermore, the authors want to indicate, that a more comprehensive consideration of the environmental impact was initially intended, with further impact indicators such as ozone depletion potential (odp), eutrophication potential (ep) and photochemical oxidant creation (pocp) being considered. unfortunately, the respective environmental data of cfrp were either not available nor valid. updates in the data availability are constantly checked and re-evaluated by the authors. 5. conclusion the intention of this paper was to evaluate, whether fibre-reinforced polymers are a possible environmentally friendly alternative to conventional steel reinforcement in concrete structures or not. in a first step, the varying fibre and matrix material types were solely evaluated per unit weight, at which carbon fibres and cfrp reinforcement showed by far the highest global warming potential. subsequently, a more detailed environmental assessment of cfrp reinforcement on material level showed that steel reinforcement is more favourable when considering solely the declared unit of 1 kg as well as the declared unit 588 vol. 33/2022 non-metallic reinforcement for concrete structures figure 8. pedestrian bridge examples, from top to bottom: b1-ccb (solidian gmbh), b2-rcb and b3-sb (mathias hammerl). figure 9. comparison of impact indicators of various pedestrian bridge types (normalized values), extracted from [5]. figure 10. comparison of the gwp of various pedestrian bridge types (normalized values) divided by the used building materials, extracted from [5]. figure 11. comparison of the ap of various pedestrian bridge types (normalized values) divided by the used building materials, extracted from [5]. figure 12. comparison of the adpf of various pedestrian bridge types (normalized values) divided by the used building materials, extracted from [5]. 589 j. anderson, m. hammerl, b. kromoser acta polytechnica ctu proceedings in combination with the tensile or rather the yield strength of the respective material. nevertheless, the consideration of the reinforcements’ performance as well as the density of the material showed a lesser environmental impact of cfrp reinforcement in comparison with conventional steel reinforcement. in a second step, an assessment and comparison of the superstructure of a carbon concrete, a reinforced concrete and a mild steel pedestrian bridge outlined the environmental benefits of the carbon concrete bridge type. the potential of cfrp as an environmentally friendly alternative to steel reinforcement in the case of pedestrian bridges could be shown. it has to be mentioned that overall conclusions of this application example to other structures are not acceptable: an individual profound analysis is always necessary. furthermore, only the production stage of the reinforcement components and pedestrian bridge materials was considered. a more holistic consideration of further life cycle stages with an extension to other structural components is highly recommended as an optimal continuation of this study. acknowledgements the authors would like to thank dirk neuburg from the municipal department 29 in vienna and dr. sergej rempel from solidian gmbh for the provision of the original bridge documents and additional information used for the bridge examples. they would like to thank clare broadbent from the world steel association for the provision of environmental data as used within the cradle-togate lca. the authors also want to thank dr. andrea hohmann for her support in the early stage of the assessment. a part of the data collection was done within the scope of the vif project 2019 "potentiale von nichtmetallischer bewehrung im infrastruktur-betonbau". the authors want to thank the customers austrian railways (öbb), the austrian autobahn and highway financing stock corporation (asfinag), the bundesministerium für klimaschutz, umwelt, energie, mobilität, innovation und technologie (bmk) as well as the austrian research promotion agency (ffg) for the management. references [1] international energy agency and the united nations environment programme. 2018 global status report: towards a zero-emission, efficient and resilient buildings and construction sector, 2018. https: //wedocs.unep.org/handle/20.500.11822/27140. [2] oecd. global material resources outlook to 2060: economic drivers and environmental consequences (paris: oecd publishing), 2019. https: //www.oecd.org/environment/waste/highlights-glo bal-material-resources-outlook-to-2060.pdf. [3] g. mauschitz.emissionen aus anlagen der österreichischen zementindustrie, 2018. [4] b. kromoser, p. preinstorfer, j. kollegger. building lightweight structures with carbon-fiber-reinforced polymer-reinforced ultra-high-performance concrete: research approach, construction materials, and conceptual design of three building components. structural concrete 20(2):730-44, 2018. https://doi.org/10.1002/suco.201700225. [5] n. stoiber, m. hammerl, b. kromoser. cradle-togate life cycle assessment of cfrp reinforcement for concrete structures: calculation basis and exemplary application. journal of cleaner production 280, 2021. https://doi.org/10.1016/j.jclepro.2020.124300. [6] cml. department of industrial ecology, cml-ia characterisation factors, 2016. https://www.universi teitleiden.nl/en/research/research-output/scie nce/cml-ia-characterisation-factors. [7] world steel association. life cycle inventory (lci) request form, 2019. https://www.worldsteel.org/steel-by-topic/lif e-cycle-thinking/lca-lciform.html. [8] world steel association. world steel in figures 2019, brussels, 2019. [9] a. hohmann. life cycle assessment of manufacturing processes for cfrp structures to identify optimization potentials systematic methodology for estimating the environmental impact of manufacturing process chains, dissertation: technical university of munich, 2018. [10] pwc. life cycle assessment of cfgf continuous filament glass fibre products, 2016. https://www.gl assfibreeurope.eu/wp-content/uploads/2016/11/lc a-report-cfgf-products_20161031_pwc.pdf. [11] k. azrague, m.r. inman, l-i. alnaes l-i, et al. life cycle assessment as a tool for resource optimisation of continuous basalt fibre production in iceland, 2016. https://dc.engconfintl.org/cgi/viewcontent.cgi ?article=1009&context=lca_waste. [12] pwc. epdla life cycle inventory of polymer dispersions, 2015. https://specialty-chemicals.eu/ wp-content/uploads/2017/07/4_epdla-life-cycle -assessment-lca-summary-report.pdf. [13] eucia. eco impact calculator for composites, 2021. https://ecocalculator.eucia.eu/account/log in?returnurl=%2f. [14] plasticseurope. eco-profiles, 2005. https://www.plasticseurope.org/de/resources. [15] l. laiblová, j. pešta, a. kumar, et al. environmental impact of textile reinforced concrete facades compared to conventional solutions lca case study. materials 12(19), 2019. https://doi.org/10.3390/ma12193194. [16] solidian. technical data sheet solidian rebar 04-cce, 2020. https://solidian.com/wp-content/u ploads/solidian-rebar-tds.pdf. [17] t. helbig, k. unterer, c. kulas, et al. fußund radwegbrücke aus carbonbeton in albstadt-ebingen. betonund stahlbetonbau 111(10):676-85, 2016. https://doi.org/10.1002/best.201600058. 590 https://wedocs.unep.org/handle/20.500.11822/27140 https://www.oecd.org/environment/waste/highlights-global-material-resources-outlook-to-2060.pdf https://doi.org/10.1002/suco.201700225 https://doi.org/10.1016/j.jclepro.2020.124300 https://www.universiteitleiden.nl/en/research/research-output/science/cml-ia-characterisation-factors https://www.worldsteel.org/steel-by-topic/life-cycle-thinking/lca-lciform.html https://www.glassfibreeurope.eu/wp-content/uploads/2016/11/lca-report-cfgf-products_20161031_pwc.pdf https://dc.engconfintl.org/cgi/viewcontent.cgi?article=1009&context=lca_waste https://specialty-chemicals.eu/wp-content/uploads/2017/07/4_epdla-life-cycle-assessment-lca-summary-report.pdf https://ecocalculator.eucia.eu/account/login?returnurl=%2f https://www.plasticseurope.org/de/resources https://doi.org/10.3390/ma12193194 https://solidian.com/wp-content/uploads/solidian-rebar-tds.pdf https://doi.org/10.1002/best.201600058 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0094 acta polytechnica ctu proceedings 34:94–97, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague detection of the presence of microorganisms during production of plasterboard hana sekavováa, b, ∗, zdeněk prošeka, george karra’ac a czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic b knauf praha spol. s r.o., mladoboleslavská 949, 197 00 prague 9, czech republic c lavaris s.r.o., areál šroubáren 43, 252 66 libčice nad vltavou, czech republic ∗ corresponding author: hana.sekavova@fsv.cvut.cz abstract. the article introduces possibility how to use special material from plasterboard back to the production of plasterboards. the recycled material was prepared from standard plasterboards knauf white (it means plasterboard type a according to european standard en 520) [1]. plasterboards were crushing, micromilled and separated from the paper on recycling line, which was developed in company lavaris s.r.o. [2]. this material was used as an inert to the standard recipe instead of stucco. there were checked the influence on properties of final plasterboards in comparison with standard plasterboards. keywords: gypsum plasterboard, stucco, recycling, inert material. 1. introduction the massive expansion of use gypsum plasterboards in czech republic started in 1990s. the first years were boards imported from neighboring countries and also came first producers and built their plants there [3]. the gypsum plasterboard plant knauf praha spol. s.r.o. is situated in počerady, direct next to power plant počerady, was opened in year 1994. the capacity of plant is millions square metres per year and the scrap together with customer´s scrap is in thousands tons per year and this amount is sent to landfills. in the world there are sent to landfills more than 15 million tons of gypsum waste from plasterboards each year [4]. the very comprehensive and clear overview of problematic recycled gypsum in eurepean union is in research presented by a.j. rivero et al. [5]. this research shows that recycling of gypsum is neccesary because of hundred thousands tons of waste, which is send to landfills, because of closing stocks of natural and also syntetic gypsum and last but not least the influence for energy consumtion and emissions. in the beginning, the fgd (flue gas desulphurization) gypsum is also kind of waste and m.a. pedrenorojas et al. [6] consider use of fgd gypsum such as use of waste. in their research is presented successfully use of fgd gypsum and waste from production from plasterboard production in plasters. the plasterboards only from recycled material was produced in laboratory condition. the recycling was with calcination, the material, which was produced, was stucco with binding properties – calcium sulphate hemihydrate [7]. there was also watched influence on properties of multiple recycled gypsum from plasterboards. it was processed, that gypsum waste is feasible until the third cycle [8]. the recyled gypsum from plasterboards can be used not only back to production of plasterboards. the essential use is also in portland cement manufacture and as a soil amendment. possibility is use this material as a compost amendment and for animal bedding [9]. the development of new sources and gypsum and recycling of gypsum products is necessary because in the near future is planned the shutdown of coal fired power plants. it means no fgd gypsum and also there is expectation that environmental impacts related to electricity consumption will decrease [10]. the target of this research is find the most safe amount of recycled material from gypsum plasterboards what is possible to use as an inert in plasterboard production. 2. material the material, which was used for this trial, was plasterboards knauf white and knauf green. knauf white is trademark for standard plasterboard type a and knauf green is trademark for standard impregnated boards type h2 [1]. these boards are produced in knauf praha spol. s r.o. in plant počerady. the plasterboards are produced of stucco, which is calcinated of fgd gypsum. the fgd gypsum is produced during the desulphurization process flue gases in power plant, where is burned the brown coal. this fgd gypsum has very high purity (96 % on average). it means there is possibility to use inert to the recipes and don´t lose the binding properties of slurry. the process of calcination (1) and hydration (2) is quite very simple and the energy, which is necessary 94 https://doi.org/10.14311/app.2022.34.0094 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 34/2022 detection of the presence of microorganisms for production of stucco, is very low in comparison with production of cement. the energy consumption for production 1 t of stucco from fgd in rotary kiln is 250–350 kcal depending on the moisture content. however, if we use the inert material instead of stucco, we can save corresponding quantity of heat energy for stucco production. caso4 · 2h2o + heat → → caso caso4 · 1/2h2o + 1/2h2o 110 − 170 ◦c → caso caso4ii + 2h2o > 200 ◦c . . . anhydrit ii (aii) → cao + so2 + o2 + caso4i > 800 ◦c . . . anhydrit i (ai) (1) caso caso4 · 1/2h2o + 1/2h2o → → caso4 · 2h2o + heat (2) the inert could be also material which could be produced during recycling of gypsum plasterboards. there is a quite big amount scrap, which is formed during the production of plasterboards and also there is a lot of scrap, which is formed during construction (customer scrap). on the other hand, the recycled gypsum can be used as a retarder during rehydration of cement [11]. figure 1. schema of production of plasterboards. schema of production of plasterboards is on the figure 1. for this topis is important point 2, where the all of components are mixed in the mixer and the gypsum slurry is put between face and back paper. the second important place is point 5. it is knife position, where the endless belt is divided to boards and where the wet scrap is formed. after dryer (point 8) are boards sawed and milled to precise dimension (point 10) and here is formed dry scrap. in year 2019 was counted that approximately 5700 tons of scrap was sent to landfill only from plasterboards produced in počerady plant (internal and external scrap). there is solved only influence of recycled material from dry knauf white boards to behavior of production and products in this paper, but there is showed inputs and outputs of recycling process in počerady plant on figure 2. the photo of raw materials and outputs of recycled proces is showed on figure 3 (first line from left: fgd gypsum, stucco, recycled paper; second line from left: recycled material 2–5 mm, recycled material 0–2 mm, recycled material 0–0.05 mm). the plasterboards knauf white were input and output from the trial too. in the beginning was taken boards, which were not fulfill all of the requirements (european and internal standards) and they usually are throw out as a scrap. this material was delivered to company lavaris s.r.o. and there was crushed, milled and it was separated out of the paper too. the process of crushing and milling was with cooling in order to avoid recalcination. output is calcium sulphate dihydrate. there is process of crushing and milling on the figure 4. in the beginning of this project was assumed prepare granulometry of inert same as used stucco (therefore the fgd gypsum – also used as an inert). on the figure 5 is a grain size curve of počerady stucco (purple one) and defined milled material (green one). the maximal size of grain was 1 mm. from the input material was prepared recycled material of three fractions: 2–5 mm, 1–2 mm and 0–1 mm. each fraction formed approximately one third of the original volume. this material (fraction 0–1 mm) was used to the production. it was prepared for the existing equipment. unfortunatelly the dosing system was not able to dose this material fluently. the very fine material made an arches and the material was very often blocked. because of this trial production was stopped and the rest of material was thoroughly mixed with the material grain size 1–2 mm. this material was dosed without any problems. for second trial was prepared material grain size 0–2 mm. there was also milled small amount of knauf green boards for verification assumption, that impregnated boards cannot be used as an inert to the gypsum slurry. 3. experimental methods and results 3.1. production the milled material was used in production in 3 variants. the first one was use recycled material with grain size 0–2 mm which was mixed from two fractions (0– 1 mm and 1–2 mm). this situation was describe in capture material. this variant has three steps (1.) standard recipe with 3 % recycled material. (2.) standard recipe without 3 % stucco with 4.5 % recycled material. (3.) standard recipe without 4 % stucco with 6 % recycled material. edit: percent means % wt. of amount of stucco in standard recipe 95 h. sekavová, z. prošek, g. karra’a acta polytechnica ctu proceedings figure 2. diagram of recycled process in počerady plant. figure 3. raw materials and outputs from the recycled process. figure 4. lavaris recycling babyline. each step took about 5 minutes. there was watched figure 5. diagram of granulometry – stucco and recycled material. only small changes on gypsum slurry. any other check didn´t notice any abnormalities. the second variant was long term trial with material with fraction 0– 2 mm. during second trial was standard recipe included inert (raw fgd), because of that was only fgd gypsum replaced with recycled material 1:1. the amount was 3 % wt. of stucco. this trial took 55 minutes and there was consumed almost 500 kg of recycled material. during the change material and trial production was not noticed any abnormalities on production line. the third one was verification assumption, that impregnated boards cannot be used as an inert to the gypsum slurry. there was used fine milled material 0–1 mm with median under 0.5 mm. there was dosed only 6 g/m2 of this material it means less than 0.002 % wt. of stucco. immediately afterwards there was a massive decline of gypsum slurry on the forming table due to degradation of foam, which is integral part of gypsum slurry. the first assumption was hereby confirmed. 96 vol. 34/2022 detection of the presence of microorganisms 3.2. output quality control output quality control consists of the visual, mechanical, and chemical tests. the set of visual controls consists of checking defects and abnormalities on the surface of boards. the boards produced during trials had no defects like bubbles or blisters under the paper. the porosity of gypsum core was without changes. mechanical properties are tested according to european standard en 520 [1]. the breaking loads was in the same range of values like boards produced with standard recipe. there was not any increase but also any decrese of values. it means all of values were between 610 and 620 n in lengthwise and between 210 and 230 n in crosswise. the most exposed and watched test is paper bonding. there is three tests of paper bonding. the first one is paper bonding direct after drying process, next one is paper bonding after 24 hours of storage and third one is paper bonding after 24 hours storage in moist box (or climatic chamber). all of the test use cross cut method, which is showed on figure 6. there is also the results of paper bonding of trial boards. there was watched no decrease of bonding between paper and gypsum core. figure 6. results of paper bonding cross test. 4. conclusions all of the trials were successful and they fulfill all of the expectation. the safe amount of recycled material is 6 % wt. of stucco. this amount of recycled material had no influence to behavior of production process and quality of product. the benefit of use inert is safe energy for production of stucco. on the other hand the recycled material is more expensive that raw fgd gypsum which is also used as an inert. important is that raw fgd gypsum is better use for production of stucco and from an ecological point of view is better use the recycled material as an inert instead of send it to the landfill. definitely this trial open the possibility to increase the amount of inert in recipes in the future. acknowledgements the research is carried out within the project mpo trio č. fv30359 “recyklace sádrokartonových desek a nová materiálová využití s přidanou hodnotou – gipsrec” (gypsum board recycling and new value-added material uses). references [1] čsn en 520+a1. sádrokartonové desky definice, požadavky a zkušební metody. praha: úřad pro technickou normalizaci, metrologii a státní zkušebnictví 2010. [2] h. sekavová, j. herrmann, z. prošek, et al. the influence of the use of recycled gypsum on the properties of gypsum products in labor conditions. acta polytechnica ctu proceedings 26:81–85, 2020. [3] m. nyč. sádrokarton. grada publishing as, 2005. [4] a. ahmed, k. ugai, t. kamei. investigation of recycled gypsum in conjunction with waste plastic trays for ground improvement. construction and building materials 25, 2011. https://doi.org/10.1016/j.conbuildmat.2010.06.036. [5] a. rivero, r. sathre, j. navarro. life cycle energy and materiál flow implications of gypsum plasterboard recycling in the european union. resources, conservation and recycling 108, 2016. https://doi.org/10.1016/j.resconrec.2016.01.014. [6] m. pedreno-rojas, i. flores-colen, j. d. brito, c. rodríguez-linán. influence of the heating proces on the use of gypsum wastes in plasters: mechanical, thermal and enviromentalanalysis. journal of cleaner production 215, 2019. https://doi.org/10.1016/j.clepro.2019.01.053. [7] a. erbs, a. nagalli, k. q. de carvalho, et al. development of plasterboard sheets exclusively from waste. journal of building engineering 44, 2019. https://doi.org/10.1016/j.jobe.2021.102524. [8] a. erbs, a. nagalli, k. q. de carvalho, et al. properties of recled gypsum from gypsum plasterboards and commercialgypsum throughout recycling cycles. journal of cleaner production 183, 2018. https://doi.org/10.1016/j.jclepro.2018.02.189. [9] čsn en iso 846. plasty hodnocení působení mikroorganismů. praha: úřad pro technickou normalizaci, metrologii a státní zkušebnictví 2019. [10] k. weimann, c. adam, m. buchert, j. sutte. enviromental evaluation of gypsum plasterboard recycling. minerals 2021. https://doi.org/10.3390/min11020101. [11] n. f. medina, f. hernández-olivares, x. arroyo, et al. characterization of a more sustainable cement produced with recycled drywall and plasterboards as set retarders. construction and building materials 124, 2016. https://doi.org/10.1016/j.conbuildmat.2016.08.140. 97 https://doi.org/10.1016/j.conbuildmat.2010.06.036 https://doi.org/10.1016/j.resconrec.2016.01.014 https://doi.org/10.1016/j.clepro.2019.01.053 https://doi.org/10.1016/j.jobe.2021.102524 https://doi.org/10.1016/j.jclepro.2018.02.189 https://doi.org/10.3390/min11020101 https://doi.org/10.1016/j.conbuildmat.2016.08.140 acta polytechnica ctu proceedings 34:94–97, 2022 1 introduction 2 material 3 experimental methods and results 3.1 production 3.2 output quality control 4 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0032 acta polytechnica ctu proceedings 34:32–37, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague microstructure reconstruction via artificial neural networks: a combination of causal and non-causal approach kryštof latkaa, ∗, martin doškářb, jan zemanb a gymmázium nový porg, pod krčským lesem 25, 142 00 prague 4, czech republic b czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: latka@novyporg.cz abstract. we investigate the applicability of artificial neural networks (anns) in reconstructing a sample image of a sponge-like microstructure. we propose to reconstruct the image by predicting the phase of the current pixel based on its causal neighbourhood, and subsequently, use a non-causal ann model to smooth out the reconstructed image as a form of post-processing. we also consider the impacts of different configurations of the ann model (e.g., the number of densely connected layers, the number of neurons in each layer, the size of both the causal and non-causal neighbourhood) on the models’ predictive abilities quantified by the discrepancy between the spatial statistics of the reference and the reconstructed sample. keywords: microstructure reconstruction, neural network, causal neighbourhood, non-causal neighbourhood. 1. introduction multi-scale modelling is a powerful predictive tool that bypasses the need for complex constitutive laws by performing auxiliary calculations at lower scales, using a characteristic sample of a material microstructure [1]. such a sample can be easily extracted when analysing a material with regular, periodic arrangement of material’s phases; in case of a material with stochastic microstructure, the representative sample is typically constructed artificially such that it matches selected spatial statistics of the material microstructure [2]. while originally the representative samples were generated via optimization approaches, e.g. [3], recently, reconstruction methods relying on machine learning have started to emerge, using various frameworks including markov random fields [4], deep adversarial neural networks [5, 6], or supervised learning using classification trees [7]. several papers include references to causal and non-causal neighbourhood which both prove to be effective ways of extracting input data for the chosen framework. however, the causal approach generally seems to be the preferred one, with one of these papers even claiming that their model cannot generate valid results if based on a non-causal neighbourhood [4]. the objective of our work is to reconstruct an image of a microstructure from an almost random noise with microscopic properties as similar as possible to the original image. while many of the aforementioned proposed approaches are implementation-complex, we present a simple method using the tensorflow framework with the keras sequential api [8]. we closely follow the methodology of bostanabad and coworkers, [7]; however, instead of using classification trees, we use two distinct artificial neural networks (anns), where the first network reconstructs the general pattern of the microstructure and the second network denoises and smoothes out the previously reconstructed pattern. for simplicity, we study only two-phase materials, i.e. we test the framework with black and white images. our implementation and data are publicly available at a gitlab repository [9]. 2. methodology the proposed reconstruction method comprises two distinct steps: (1.) the reconstruction of a general shape of the microstructure from a random noise with margins of the reference image used as a seed. (2.) the smoothing procedure of the reconstructed image that improves the local features of the microstructural geometry. 2.1. reconstructing microstructural geometry in step 1, the overall material distribution should be outlined in the reconstructed sample, rendering the key features of the microstructural geometry. we opted for a sequential approach, in which new values of individual pixels in the discrete, pixel-like representation of the newly generated microstructural geometry are predicted from the values previously determined at antecedent positions (thus the causal approach), because such an approach has already proven its merits in microstructural reconstruction, cf. [7]. 32 https://doi.org/10.14311/app.2022.34.0032 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 34/2022 microstructure reconstruction via ann hr figure 1. illustration of the causal neighbourhood with hr = 2 pixels around the central pixel highlighted in dark grey. the red pixels represent the input data for the ann trained in step 1; however, in order to extract an input structure of a rectangular shape, the black and yellow pixels are also included in the inputs but their values are discarded and replaced by random binary values. to this end, we define a rectangular causal neighbourhood of (2hr + 1)×(hr + 1) pixels with parameter hr being the given neighbourhood radius; see fig. 1 for an illustration. note that the positions highlighted in red constitute the real input data; the dark grey and yellow pixels are only a padding (without value) such that the whole input features a regular 2d shape and, consequently, pooling layers can be easily applied. the actual value of the dark grey pixel serves as a label during an extraction of the training data from a reference image. the neural network for predicting the pixel values in step 1 is designed such that the rectangular input is first subsampled using a nrp ×nrp pooling layer, flattened and passed through several densely connected layers. based on our numerical experiments, the maximum pooling layer consistently delivered better results than the average pooling layer. the setup of the training procedure and the effect of the remaining network parameters, namely the number of densely connected layers nr`, the number of neurons in each layer n r n, the size of the pooling layer nrp and the neighbourhood radius hr, are discussed later in section 4. once the network is trained, a new microstructural realization is generated by iterating over pixels of a to-be-reconstructed image in a raster scan order. consequently, an initial microstructural geometry must be provided in a margin of width hr at the left, top, and bottom edge of the image; the initial values in the remaining part of the image are irrelevant and we generate them as a random binary noise; see fig. 2b. 2.2. smoothing procedure outputs of the model trained in step 1 usually contain the main features of the trained microstructural geometry; however, local details are typically polluted by random noise; compare figs. 2a and 2c. for step 2 we train an additional neural network to smooth out the image and correct irregularities in the image generated in the step 1. this time, the model works with a complete, i.e. non-causal, square neighbourhood around the central pixel (illustrated in figure 3), usually of a smaller neighbourhood radius hs than in the causal model in step 1. again, the two-dimensional structure of the input is needed to facilitate a subsampling/pooling layer, which is then followed by flattening and passing through two densely connected layers. the impact of the actual choice (i.e. average vs maximum pooling) of the subsampling layer is discussed in section 4). to increase robustness of the trained model and prevent if from learning to simply copy the value of the central pixel, we introduce two errors: (i) the value of the dark grey pixel, which is used as a label in the training, is always randomized in the inputs, and (ii) we also randomly choose value for ξ × 100% of red pixels from fig. 3. 3. error quantification in order to assess the performance of the proposed models beyond a visual inspection, we compare generated microstructural samples to the reference images in terms of spatial statistics. the most straightforward spatial statistics is a volume fraction φ of a chosen phase (the white phase, i.e. pixels with value 1, in our case). we define error εφ as an absolute value of the difference between the volume fraction φref in the reference sample and the volume fraction φgen in the generated microstructure, εφ = |φref −φgen| . (1) the second spatial statistics considered in our work is the two-point probability function s2(x), which states the probability of finding two points separated by x in a given phase. since all our data are represented as a regular grid of values, the discrete version of s2(x) can be easily computed using the fast fourier transform [2]. consequently, we quantify the discrepancy in the reference and a generated microstructure by means of their discrete two-point probability functions sref2 ∈ rni×nj and s gen 2 ∈ r ni×nj as εs2 = ‖sgen2 − s ref 2 ‖f ninj , (2) where ‖a‖f is the frobenius matrix norm [10] ‖a‖2f = ni∑ i=1 nj∑ j=1 (ai,j)2 . (3) finally, to quantify the effect of our smoothing noncausal model, we add the third error metric εd that captures the level of local heterogeneity. assuming a two-phase medium, a microstructure can be represented with a boolean matrix m ∈{0, 1}ni×nj . for each pixel we can computed a local quantity di,j as a sum of the averaged absolute differences between the 33 k. latka, m. doškář, j. zeman acta polytechnica ctu proceedings (a) (b) (c) figure 2. microstructure reconstruction process using the causal model. after training the neural network model on input data from the initial microstructure (a), the trained model is used to sequentially predict pixel values in a raster scan order using previously predicted pixel values from the causal neighbourhood. hs figure 3. illustration of the non-causal neighbourhood of a neighbourhood radius of hs = 2 pixels. the black pixel represents the central pixel and the red pixels represent non-causal neighbourhood. when extracting the neighbourhood, the central black pixel is represented as a random binary value. value of the central pixel and its neighbouring eight pixels, di,j = 1 8 1∑ k,l=−1 |mi+k,j+l −mi,j| . (4) the error εd is then computed again as an average over the image excluding the one-pixel wide margin, i.e. εd = 1 (ni − 2)(nj − 2) ni−1∑ i=2 nj−1∑ j=2 di,j . (5) the reasoning behind this error measure is that if we consider the phases of pixels in a very small neighbourhood around the central pixel, the number of pixels whose phase is different to that of the central pixel will be lower if the edges are properly smoothed out. even though this might not necessarily be true for the pixels which form the edge of the reconstructed pattern (and thus, the black and white phase must switch), it will apply on a larger scale (hence, we compute the sum of the values of dij, for all pixels in the image). therefore, in theory, the lower the value of εd is for an image, the more smoothed out the image should be. 4. results we report the effect of parameters on the quality of the microstructural reconstruction quantified with the error measures introduced in the previous chapter. first, we focus on parameters of the reconstruction model. tables 1 and 2 illustrate the impact of altering number of neurons nrn 9 16 25 nr` = 2 0.046 0.031 0.036 nr` = 3 0.031 0.049 0.029 nr` = 4 0.002 0.010 0.021 table 1. values of the volume fraction error εφ, depending on the number of layers nr` and the number of neurons in each layer nrn. number of neurons nrn 9 16 25 nr` = 2 1.36×10−4 9.36×10−5 1.05×10−4 nr` = 3 9.32×10−5 1.46×10−4 7.57×10−5 nr` = 4 2.96×10−5 3.71×10−5 6.70×10−5 table 2. values of the two-point probability error εs2 , depending on the number of layers nr` and the number of neurons in each layer nrn. the number of densely connected layers nr` and the number of neurons in each layer nrn, while keeping the neighbourhood radius hr and the pooling size nrp constant, on the volume fraction error εφ and the twopoint probability error εs2 , respectively. in particular, we set hr = 15 and nrp = 3 as these values produced the most visually appropriate reconstructions in early tests of the model. the next two tables, tabs. 3 and 4, summarize the sensitivity study investigating the influence of the neighbourhood radius hr and the pooling size nrp on the reconstruction. this time, all the tests were performed with nr` = 2 densely connected layers with nrn = 16 neurons in each layer. 34 vol. 34/2022 microstructure reconstruction via ann (a) (b) figure 4. effect of the smoothing network based on the non-causal model: (a) reconstructed image serving as an input, (b) microstructure corrected by the smoothing model. both images are of size 200 x 200 pixels. neighbourhood radius hr 10 12 15 nrp = 2 0.491 0.298 0.006 nrp = 3 0.058 0.046 0.031 table 3. values of the volume fraction error εφ, depending on neighbourhood radius hr and pooling size nrp. neighbourhood radius hr 10 12 15 nrp = 2 2.50×10−3 1.24×10−3 4.05×10−5 nrp = 3 1.81×10−4 1.36×10−4 9.36×10−5 table 4. values of the two-point probability error εs2 , depending on neighbourhood radius hr and pooling size nrp. the next three tables, i.e. tables 5, 6, and 7, summarize the parametric study for the smoothing model with non-causal neighbourhood. we chose the reconstructed image obtained by the first model using nr` = 2 layers, each with n r n = 16 neurons, a neighbourhood radius of hr = 15 and a pooling size of nrp = 3 as an input to the model and compared the resulting smoothed-out image to the original microstructure (before reconstruction), recall fig. 2a, in terms of the error measures introduced in section 4. we carried out two sets of test: one for the average and one for the maximum pooling 2d layer. in each set, we altered the radius of the non-causal neighbourhood hs as well as the magnitude of the artificially introduced noise ξ. each set rendered three tables as we inspected also the level of local heterogeneity εd, in addition to the errors εφ and εs2 already reported for the generative model. 5. discussion first, we add an observation regarding the three error measures defined in section 3 and our visual perception of the reconstructed microstructures. the first error measure, εφ, served as a coarse check that the volume fraction in the reconstructed image is similar neighbourhood radius hs 5 7 10 m ax ξ = 0.05 0.041 0.002 0.040 ξ = 0.10 0.004 0.036 0.033 ξ = 0.15 0.051 0.027 0.033 av er ag e ξ = 0.05 0.058 0.000 0.033 ξ = 0.10 0.037 0.003 0.030 ξ = 0.15 0.044 0.029 0.036 table 5. values of the volume fraction error εφ, depending on the magnitude of artificially introduce noise ξ, neighbourhood radius hs, and the type of pooling layer used (max or average). neighbourhood radius hs 5 7 10 m ax ξ = 0.05 1.22×10−4 3.15×10−5 1.20×10−4 ξ = 0.10 3.24×10−5 1.06×10−4 9.88×10−5 ξ = 0.15 1.23×10−4 7.21×10−5 9.88×10−5 av er ag e ξ = 0.05 1.76×10−4 3.08×10−5 1.01×10−4 ξ = 0.10 1.12×10−4 3.25×10−5 8.98×10−5 ξ = 0.15 1.31×10−4 8.90×10−5 1.08×10−4 table 6. values of the two-point probability error εs2 , depending on the magnitude of artificially introduce noise ξ, neighbourhood radius hs, and the type of pooling layer used (max or average). to the original; however, it could not assess how similar the reconstructed pattern is to the original. for this purpose, we adopted εs2 , based on the two-point correlation function, as we excepted it ot be better suited to compare the reconstructed pattern to the original. nevertheless, we noticed a significant discrepancy between the values of εs2 for each model and the visual similarity of the reconstructed pattern to the original one. for example, the generative model using nr` = 2 layers, each with n r n = 16 neurons, neighbourhood radius of hr = 15 and a pooling size of nrp = 3, could probably be considered as the most accurate in terms of the visual pattern (the image reconstructed using this model is in figure 2c), but only seventh 35 k. latka, m. doškář, j. zeman acta polytechnica ctu proceedings neighbourhood radius hs 5 7 10 m ax ξ = 0.05 0.130 0.133 0.123 ξ = 0.10 0.135 0.129 0.124 ξ = 0.15 0.123 0.130 0.127 av er ag e ξ = 0.05 0.113 0.120 0.112 ξ = 0.10 0.114 0.119 0.112 ξ = 0.15 0.114 0.115 0.113 table 7. values of the local heterogeneity error εd, depending on the magnitude of artificially introduce noise ξ, neighbourhood radius hs, and the type of pooling layer used (max or average). for comparison, the value of εd for the reference image is 0.097. best according to εs2 . we conclude that other spatial statistics such as two-point cluster function or lineal path should be added to the suite of error measures as well to capture both the global distribution of a microstructure and its local characteristics. on the other hand, the assessment of the smoothing model by εd values was generally in accordance with the quality of the visual appearance of the reconstructed images. our results show that taking more layers with less neurons in each was favourable to the approach with less but more populated layers. perhaps as expected, the larger neighbourhood radius was considered during training of the generative model, the better the results were. surprisingly, nrp = 2 pooling size was better for the largest neighbourhood radius, while larger pooling was preferential in all other cases. in the case of the smoothing model, considering larger neighbourhood radius hs beyond certain threshold (in our case hs = 7) did not improve its performance. this can be attributed to the different purpose of both models; while the generative model needs information from distant points to properly distribute the material with the sample, the smoothing model is by its nature local. on the other hand, larger pooling layer was consistently outperforming the smaller one in all tests. most importantly, we noticed that the values of εd were significantly lower for the smoothing models using an average pooling layer instead of a maximum pooling layer. this is probably due to the fact that the average pooling layer ignores sharp features in the image (e. g. individual pixels whose phase was not correctly identified in the part 1 of the reconstruction), which allowed us to smooth out the edges of the reconstructed pattern in the image. however, it is important to emphasize that these observations are specific for the considered microstructure. 6. conclusions despite the simplicity of the proposed ann-based model, accompanied by the ease of implementation facilitated by the tensorflow framework and the keras sequential api, the model generates meaningful microstructural geometries. the combination of a causal model used for reconstruction and a noncausal smoothing model in particular yielded satisfying results, cf. figure 4, considering the fact that the models knew only a limited local information. we believe that even better results can be obtained by, e.g., incorporating the considered errors directly in the loss function during the training process of individual networks. yet, this remains to be done in our future work. the need for the margin of initial values during reconstruction might be seen as a limitation restricting the model to generating microstructural samples only as large as the reference one, from which the margin can be easily copied. however, a possible solution is to take the reference sample, dismember it into pieces and reorder the pieces so that they form the margin of desired size. alternatively, starting from different parts of the reference microstructure, a set of smaller samples can be generated; these samples can be then assembled together while blending the microstructure in their overlaps using, e.g., image quilting [6]. acknowledgements m. doškář and j. zeman gratefully acknowledge support by the czech science foundation, project no. 19-26143x. references [1] m. g. d. geers, v. g. kouznetsova, k. matouš, j. yvonnet. homogenization methods and multiscale modeling: nonlinear problems. in e. stein, r. de borst, t. j. r. hughes (eds.), encyclopedia of computational mechanics second edition, pp. 1–34. john wiley & sons, ltd, chichester, uk, 2017. https://doi.org/10.1002/9781119176817.ecm107. [2] s. torquato. random heterogeneous materials: microstructure and macroscopic properties. no. 16 in interdisciplinary applied mathematics. springer, new york, 2002. [3] j. zeman, m. šejnoha. from random microstructures to representative volume elements. modelling and simulation in materials science and engineering 15(4):s325–s335, 2007. https://doi.org/10.1088/0965-0393/15/4/s01. [4] l.-y. wei, m. levoy. fast texture synthesis using tree-structured vector quantization. in proceedings of the 27th annual conference on computer graphics and interactive techniques siggraph '00. acm press, 2000. https://doi.org/10.1145/344779.345009. [5] z. yang, x. li, l. c. brinson, et al. microstructural materials design via deep adversarial learning methodology. journal of mechanical design 140(11), 2018. https://doi.org/10.1115/1.4041371. [6] d. fokina, e. muravleva, g. ovchinnikov, i. oseledets. microstructure synthesis using style-based generative adversarial networks. physical review e 101(4), 2020. https://doi.org/10.1103/physreve.101.043308. 36 https://doi.org/10.1002/9781119176817.ecm107 https://doi.org/10.1088/0965-0393/15/4/s01 https://doi.org/10.1145/344779.345009 https://doi.org/10.1115/1.4041371 https://doi.org/10.1103/physreve.101.043308 vol. 34/2022 microstructure reconstruction via ann [7] r. bostanabad, a. t. bui, w. xie, et al. stochastic microstructure characterization and reconstruction via supervised learning. acta materialia 103:89–102, 2016. https://doi.org/10.1016/j.actamat.2015.09.044. [8] m. abadi, a. agarwal, p. barham, et al. tensorflow: large-scale machine learning on heterogeneous systems, 2015. software available from tensorflow.org. [9] k. latka, m. folwarczny, m. doškář, j. zeman. aibased microstructure reconstruction. https://gitlab. com/martindoskar/ai-based-reconstruction, 2021. [10] g. h. golub, c. f. van loan. matrix computations. the johns hopkins university press, baltimore, 4th edn., 2013. 37 https://doi.org/10.1016/j.actamat.2015.09.044 https://gitlab.com/martindoskar/ai-based-reconstruction https://gitlab.com/martindoskar/ai-based-reconstruction acta polytechnica ctu proceedings 34:32–37, 2022 1 introduction 2 methodology 2.1 reconstructing microstructural geometry 2.2 smoothing procedure 3 error quantification 4 results 5 discussion 6 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.37.0054 acta polytechnica ctu proceedings 37:54–57, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague reactor criticals and spent fuel systems similarities matěj šikla, b a czech technical university in prague, faculty of nuclear sciences and physical engineerings, department of nuclear reactors, v holešovičkách 2, 180 00 prague 8, czech republic b nri řež, a. s., department of reactor physics, hlavní 130, 250 68 husinec – řež, czech republic correspondence: matej.sikl@ujv.cz abstract. many conservative assumptions for subcriticality assessment calculations have to be made in current calculations of spent fuel storage systems. this is due to a fact that there is no available database of benchmarks containing spent fuel or materials with isotopic composition like spent fuel. however one potential source of spent fuel system experiments is being omitted – there are hundreds of commercial reactors around the world containing spent fuel on some level of burnup. these reactors have to be tested for safety reasons in the beginning of each cycle and parameters of these test states are well known. in this paper a possibility of using commercial reactor critical states for code validations and for subcriticality assessments of storage systems is discussed and potential approach for criticality safety analysis is given. keywords: reactor criticals, spent fuel storage systems, calculation similarity, criticality safety analysis. 1. introduction spent fuel storage systems subcriticality limit of kef f < 0.95 is set in current czech [1] and many other countries legislation. this limit is usually fulfilled, even though some adjustments and estimations not reflecting real state have to be used when creating storage models. firstly, due to a lack of credible benchmarks with spent fuel isotopes, currently there is low knowledge about spent fuel model calculations precision and therefore reliable calculations with fresh fuel are usually performed. this also means that calculated kef f s are usually higher than reality, because by using of fresh fuel, falling trend of fuel reactivity with raising level of burnup is omitted. secondly, some conservative assumptions have to be used about neutron absorbers included in system. but if these would be too conservative, not all systems would fulfil the legislative limit. because of these adjustments, calculations for proving subcriticality of spent fuel storage system are performed on models, which are not perfectly alike to the reality of storages. subcriticality calculations of storage systems (and even criticality calculations of systems including spent fuel) would be more reliable, if a calculation code and process could be validated on spent fuel experiments. luckily one possible source of experiments is available – physical and safety start-up tests from beginning of each cycle of commercial reactors in nuclear power plants. states from these tests has sufficiently known parameters to become benchmark experiments, including mainly exactly known kef f = 1 (that is why these states are being referred as "reactor criticals") and precisely measured core parameters from instrumentation and control system. as was said, there is potential for development of more trustworthy process for spent fuel systems calculations validation. but the first milestone is to verify that reactor criticals and spent fuel systems are sufficiently similar. this is what will be discussed in this paper. 2. current approach to subcriticality assessment as was already written in the introduction, spent fuel storage systems subcriticality assessments have to be performed based on calculations using fresh fuel isotopic composition and conservative assumptions. brief summary of fresh fuel systems experiments available for code validations and generally used conservative adjustments will be given. 2.1. generally available benchmarks for code validation in an ideal case calculation codes should be validated using benchmarks of same type of systems calculations like the newly projected system it will be applied to. in reality this is impossible for spent fuel storage systems, because benchmarks with spent fuel composition are usually not described in available databases. few artificially prepared experiments with spent fuel 54 https://doi.org/10.14311/app.2022.37.0054 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 37/2022 reactor criticals and spent fuel systems similarities isotopes were described, but these usually has only units of required isotopes and not in atomic density quantities which are typically present in real spent fuel. therefore validation of calculation codes for any type of calculation is usually performed using fresh fuel experiments only. this approach is not ideal and in case of using it, calculations of models of different types should be penalized. main source of benchmarks cases with fresh fuel which can be used for validation of calculation codes is international handbook of evaluated criticality safety benchmark experiments (produced by icsbep) [2]. this database contains sets of various cases of different systems from just uranium sphere in moderator to simplified minicores (for example vver reactors type largest available minicore is with 17 fuel assemblies). isotopic composition for these benchmarks is fresh fuel and closest to real cores of nuclear reactors are systems with fresh mox fuel, because of pu included in it (from comparisons in [3]). summarized from previous text – benchmarks currently used for code validations are not similar to storage systems on sufficient level and therefore not appropriate for spent fuel systems calculations validation. 2.2. conservative adjustment of models fresh fuel isotopic composition is used for currently performed spent fuel storage subcriticality assessment. for that case there are two main generally used adjustments: • omitting of burnable absorbers, • partial boron credit. in newly produced fuel assemblies burnable absorbers (for vver it is gadolinium in gd203 mixed to some pins in fuel assembly) are used for flattening of neutron flux profile in fuel assemblies during reactor operation. these absorbers reduces reactivity and therefore have to be removed from model. technically after few cycles in reactor, their influence is negligible, but since fresh fuel is used in calculation, these absorbers have to be removed to stay conservative. in spent fuel storage pools a boric acid is dissolved to reduce reactivity. conservatively the boric acid should be completely removed from model, since it reduces reactivity. on the other side, with zero concentration of the boric acid and fresh fuel isotopic composition, kef f of the pool loaded with currently used fuel assemblies from vver–440 cycles from last years would not fulfil 0.95 kef f limit, because of their initial enrichment. rule of partial boron credit means utilization of 1/3 of a nominal boron concentration in storage models and was permitted by state supervisor. with that amount of boron in the pool system, currently used fuel with original enrichment 4.38 % meets limit requirements [4]. 3. theory of similarities as was often said in previous text, fresh fuel and spent fuel systems are not similar, but it was not said, how similarity can be measured. this section will quantify similarity. for purposes of this paper when talking about the similarity, the similarity of calculations between modelled systems is meant. when a model is prepared and calculations are performed, final kef f is based on a geometry of model, an isotopic composition and nuclear data. when nuclear data changes (each time individual value from library), it affects kef f – value describing size of this change is called "sensitivity" (of kef f to variance of nuclear data). set of this values describes computational behaving of calculation. for similarity assessment sensitivities for each isotope, reaction and energy group are needed. set of these sensitivity profiles is unique for each calculation and if compared with another systems profiles it gives information about how similar these two systems are. scale code system [5] includes module tsunami–ip, which is used for comparison of systems based on sensitivities. there are few similarity values produced by tsunami–ip comparison, but only similarity coefficient ck will be used in this paper, because it is most complex coefficient of all coefficients available. ck coefficient describes a total similarity between two systems. this value is normalized, therefore 1 means a total computational similarity, 0 means no similarity. generally value of 0.95 is considered to be sufficient for similarity , but based on [6] ck > 0.8 will be sufficient, if more than 25 experiments with this level of similarity will be available for a comparison and following (sub)criticality assessment (which will not be described in this paper). experiments with ck < 0.7 are considered to be "highly dissimilar" based on [7]. main objective of using a similarity comparison and filtering of experiments dissimilar to newly projected system is that validation of code for final calculations of new system (here for example spent fuel system) should not be performed based on dissimilar experiments or benchmarks. 4. reactor criticals states start–up tests on commercial reactors have to be performed in the beginning of each cycle after loading of new fuel. timepoint in the beginning of cycle before powering up means that there is no xe and other short term decay fission products in core, which would bring in additional uncertainty, because they decay during a shutdown when refuelling. during start–up tests, some of states are so called "hot zero power" (which implies, that power in core is considered to be negligible) and some of these start–up tests are critical (kef f = 1, from there comes reactor "criticals") – these states (on figure 1) can be used as experiments and were used in works described in this paper. 55 matěj šikl acta polytechnica ctu proceedings figure 1. reactor critical fullcore model for reactor criticals all parameters of core are known from instrumentation and control systems. therefore reactor criticals are experiments on reliability level of benchmarks, just with no detailed description in generally available database. data for reactor criticals models used for work described in this paper were taken from a database of operation histories of czech commercial reactors in dukovany and temelin, which is stored in department of reactor physics, nri rez near prague. this database contains more than 120 reactor critical states from beginnings of cycles with average core burnups within 16 to 26 gwd/tu. 5. spent fuel storage systems typical spent fuel storage systems are of two types – casks and pools. in this paper, a vver–440 storage pool filled with gd2mp with initial enrichment 4.38 % and a vver–1000 cask castor filled with tvsa–t c44e12 fuel are used for comparison. for calculations of both models are used states when spent fuel in various levels of burnup is loaded and model is filled with water with no boric acid. these states are used conservatively, because these are potentially most reactive states during lifecycle of storage’s intended usage. visualisations of spent fuel storage models are showed in figures 2 and 3. 6. software used in calculations comparison calculations were performed in tsunami–ip module of scale mentioned above, version used was 6.2.3. sensitivities were calculated in serpent 2.1.32, because during works it was found that scale code (namely tsunami–3d–k6) was not figure 2. spent fuel storage pool figure 3. spent fuel storage cask castor able to easily calculate such complex system as a fullcore model of a reactor core. calculations in serpent 2 can produce sensitivity data, but they are written in format incompatible to scale. therefore a software for preparation of sensitivities, which can read sensitivity data from serpent 2 outputs and prepare sensitivity inputs for tsunami–ip comparison calculations, had to be created. 7. results of similarity comparison for this work 19 reactor criticals states from various cycles of dukovany and temelin nuclear power plants start–up tests were taken for sensitivity calculation. storage systems were prepared artificially, with exactly given burnups on various levels. similarity coefficients were calculated in tsunami–ip and amounts of reactor criticals which are more similar to individual spent fuel storage systems than various ck values can be seen in tables 1 and 2. 56 vol. 37/2022 reactor criticals and spent fuel systems similarities fuel burnup pool required ck > [gwd/tu] kef f 0.8 0.9 0.95 10 0.890 19 18 16 15 0.862 19 19 16 20 0.836 19 19 19 25 0.814 19 19 19 45 0.720 19 6 0 table 1. amounts of reactor criticals with ck similarity higher than 0.8, 0.9 and 0.95 to pool storage systems in various levels of burnup. from table 1 can be seen, that for the storage pool there are plenty of sufficiently similar available reactor criticals experiments. in future subcriticality assessments of spent fuel storage pool, it will be possible to validate calculation codes based on reactor criticals states. fuel burnup castor required ck > [gwd/tu] kef f 0.7 0.75 0.8 5 0.801 5 1 0 15 0.796 9 4 0 25 0.744 13 6 0 45 0.649 10 4 2 table 2. amounts of reactor criticals with ck similarity higher than 0.7, 0.75 and 0.8 to storage cask castor systems in various levels of burnup on the other side castor (as can be seen from table 2) has less sufficiently similar systems. that could be for example due to a pitch between fuel assemblies or support construction itself. current results for these sets of castor and reactor criticals models are not similar enough for reliable calculations for subcriticality assessment. anyhow by comparisons presented in this section is shown that calculating states from the reactor criticals database is possible and that they are more similar to spent fuel systems than usually used fresh fuel systems are. 8. conclusions as was described in the beginning of this paper, current approach to calculations of spent fuel systems is not considered to be optimal. conservative assumptions have to be made, using these it is not automatic to fulfil legislative limits, and credit of burnup lowering reactivity is not used. non-using of burnup credit is mainly due to fact, that there is a lack of spent fuel systems for validation. in this paper the lack is resolved by using reactor critical states from start–up tests of commercial nuclear reactors. similarities between reactor criticals and spent fuel storage systems were compared using tsunami–ip module from scale code system. based on this comparison it was shown that code validated using reactor criticals would be eligible for calculations of storage pools, but not so adequate for castor cask calculations. even though similarities were high enough to storage pool only, possibility of validation of code for spent fuel systems calculation was proved when reactor criticals were used. generally it was shown and is suggested that codes can and should be validated based not only using fresh fuel benchmarks, but also using reactor criticals states. acknowledgements work presented in this report was supported by tačr grant tk02010136. references [1] vyhláška č. 329/2017 sb. o požadavcích na projekt jaderného zařízení, 2017. [2] international handbook of evaluated criciality safety benchmark experiments. nea/nsc/doc(95)03, oecd, nea no. 7382, 2016. [3] m. šikl. stanovení neurčitosti ve výpočtech kritičnosti s využitím metodiky tsunami. master’s thesis, department of nuclear reactors, fnspe, ctu in prague, 2018. http://hdl.handle.net/10467/77139. [4] m. gren, l. marková. průkaz podkritičnosti dolní kompaktní mříže bazénu edu s implementací pbc při zavezení palivem gd-2m. ujv z – 3732, 2013. [5] b. rearden, m. jessee, et al. scale code system, ornl/tm-2005/39, version 6.2.2. oak ridge national laboratory, oak ridge, tennessee, 2017. [6] b. broadhead, b. rearden, c. hopper, et al. sensitivityand uncertainty-based criticalitz safety validation techniques. nuclear science and engineering: 146, pp. 340–366, 2004, https://doi.org/10.13182/nse03-2. [7] b. rearden, m. williams, m. jessee, et al. sensitivity and uncertaintz analysis capabilities and data in scale. nuclear technology, 174:2, pp. 236–288, 2011, https://doi.org/10.13182/nt174-236. 57 http://hdl.handle.net/10467/77139 https://doi.org/10.13182/nse03-2 https://doi.org/10.13182/nt174-236 acta polytechnica ctu proceedings 37:54–57, 2022 1 introduction 2 current approach to subcriticality assessment 2.1 generally available benchmarks for code validation 2.2 conservative adjustment of models 3 theory of similarities 4 reactor criticals states 5 spent fuel storage systems 6 software used in calculations 7 results of similarity comparison 8 conclusions acknowledgements references doi:10.14311/app.2021.30.0093 acta polytechnica ctu proceedings 30:93–97, 2021 © czech technical university in prague, 2021 available online at http://ojs.cvut.cz/ojs/index.php/app detection of the presence of microorganisms during production of plasterboard hana sekavováa, b, ∗, zdeněk prošeka, pavla ryparováa a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b knauf praha spol. s r.o., mladoboleslavská 949, 197 00 prague 9, czech republic ∗ corresponding author: hana.sekavova@fsv.cvut.cz abstract. this article is focused on detection microorganisms appearing during production of gypsum plasterboards. the research is carried out within the project tačr zéta tj02000235 sádrokarton se zvýšenou odolností proti bilogické degradaci (gypsum plasterboard with increased biodegradation resistance). this article shows first part of this issue what was detection of the most critical place of production, where is the highest risk of bacterial contamination and it open the area to find suitable modification of production process or recipes of gypsum plasterboards. keywords: gypsum plasterboard, microorganism, mold, stucco. 1. introduction gypsum plasterboard is basically a composite material that consists primarily of a gypsum core and paper. it is the most widely used element of dry construction, which spread in europe in the 1970s. there was a massive expansion of this type of construction in the early 1990s in the czech republic, when imported gypsum boards from neighboring countries first began to appear on the market, and in the mid-1990s, gypsum boards produced domestically [1]. the term dry construction, of course, includes not only plasterboard, but also other types of building boards and they of course include accessories for their assembly profiles, screws, sealants, etc. the research of detection of the presence of microorganisms during production of plasterboard has two main targets. the first one is find solution for presence molds during storage, construction and use and the second one is prevention of mold during storage of waste before recycling and its subsequent introduction back into production process. recycling in general has been a much-discussed and supported topic with regard to sustainable development in recent years. research in the field of concrete recycling is very popular, but even gypsumbased building elements are not left behind. the recycling of gypsum itself is, by its nature, simple and repeatable. it is only a conversion of calcium sulphate dihydrate to its hemihydrate and back again. recyclability in the case of gypsum plasterboard is partially limited by the carton and other components added to the gypsum slurry, which reduce the purity of the calcium sulphate and some fundamentally affect the rehydration behavior. the information about the effect of the use of recycled material based on gypsum on the properties are mentioned in tasks from years 2018 and 2019 [2, 3]. influence of living organism to buildings, therefor building materials has many aspects. what can we observe at first sight is esthetic damage, but there is also mechanical damage and health effect [4]. that is the reason why there is lot of researches focus on this topic in many researches. the influence of the microoraganisms what occupied the building parts is very time solved topic. from many of them is popular detection and solution for etics [5, 6] and last but not least problematic with organism in wood construction [7]. with plasterboards have these topics similar base – stucco and gypsum like binder and paper like a product of wood. the target of this research is find the most critical place of production, where is the highest risk of bacterial contamination. 2. material at the beginning of this research was detected places and materials across the plant počerady. the schema of production of plasterboards is showed on figure 1. place no.1 is unwinding of bottom paper (face side paper). no.3 is unwinding of upper paper (backside paper). position no.2 shows main mixer preceded by preparatory station of dry and wet component. all components are mixed in main mixer and gypsum slurry is put between bottom and upper paper and this slurry together with paper continues through forming table to the setting (no.4). the gypsum slurry hardens on the setting belt. the important feature of this wet belt of product is possibility to cut. the cut is made on knife (no.5), where is cut the endless belt of plasterboard product on the individually boards. the boards are completed and transport by cross conveyor and turner (no.6) to tipple (no.7), which distributes the boards to dryer (no.8). the dryer has 8 floors and standard boards are dried in 93 http://dx.doi.org/10.14311/app.2021.30.0093 http://ojs.cvut.cz/ojs/index.php/app h. sekavová, z. prošek, p. ryparová acta polytechnica ctu proceedings figure 1. schema of production of plasterboards. sample identified microorganisms paper pulp sphingomonas paucimobilis, pseudomonas aeruginosa process water pseudomonas aeruginosa, penicillium sp. gypsum plasterboard stenotrophomonas maltophilia, penicillium sp., mucor stucco penicillium sp. table 1. identification of the samples used to determine the resistance of the materials. two lines. počerady plant has dryer which is consist of prezone, 1st zone, 2nd zone and the cooling section. during way through the dryer is evaporated all of not necessary water. the temperature of gypsum core has to be under 100 °c all the time. boards are transported after dryer by tipple (no. 9) to turner and they are completed on turner and they continue to bundler (no. 10), where are boards cut (possibly mill) to final dimension (beveling). the last process is assembly of plasterboards on transport pallets (no.11).the boards then travel to the finished goods warehouse the samples of input raw materials and final products were collected in the period from march to june 2019. the sampled plasterboards were two main products knauf white and knauf green. knauf white is the gypsum plasterboard type a. it is plasterboard with a face to which suitable gypsum plasters or decoration may be applied. knauf green is the gypsum plasterboard type h. it is plasterboard with reduced water absorption rate. these boards are produced with additeves to reduce the water absorption rate and they may be suitable for special applications in which reduce water absorption properties are required to improve the performance of the board [8]. knauf green has total water absorption under 10 % and it fulfill the category h2 according to eurepean standard en 520. the main products for production of plasterboards are stucco, water and paper. stucco is calcinated of fgd gypsum, which is produced during desulphurisation process in power plant. process water is filtrated surface water. input raw materials, what were sampled, were stucco, fgd gypsum, process water, paper pulp and paper. intermediate were gypsum slurry (after the mixer) and wet board (cut at knife, before drying process). 3. experimental methods and results 3.1. determination of a consortium of microorganisms identification of microorganisms was performed microscopically according to significant features on suitable media and according to their enzyme properties using commercial kits. enriched medium total agar (oxoid) was used to identify bacteria. cultivation was performed at 30 °c for 48 hours. molds were identifiable after culturing on czapek dox agar (oxoid) at 25 °c for three to seven days. similar microorganisms occur throughout the production process and therefore a specific source of contamination cannot be clearly identified. the most likely source of mold is the free flow of air and the movement of people and vehicles between the outdoor area and the production hall. the largest source of bacteria is pulp and process water. similar microorganisms were found in the air 94 vol. 30/2021 detection of the presence of microorganisms in plasterboard figure 2. identification of naturally occurring molds. sample kind of sample 1 gypsum plasterboard type h with paper 2 gypsum plasterboard type h without paper 3 gypsum plasterboard type a with paper 4 gypsum plasterboard type a without paper 5 gypsum slurry (hardened) type h 6 gypsum slurry (hardened) type a 7 gypsum labor slurry (stucco + water; hardened) 8 paper green 9 paper white table 2. identification of the samples used to determine the resistance of the materials. and on the surface of the production line, but their amount is negligible. results are showed in table 1. the designation of the samples in the microbiological study is given in the following table 2. individual materials were contaminated with molds of the genus penicillium, aspergillus and mucor. these molds were also found on the samples taken in production process. the highest contamination was detected in materials with paper (sample 1 and sample 3). the figure 2 shows identification of naturally occurring molds. 3.2. monitoring material degradation microbial degradation was performed according to the standard iso en 846 [9] that determines the degree of resistance of the material to the occurrence of molds (determines their resistance to biodegradation). the samples were 1 x 1 cm in size, their height depended on type of material. sterilization was performed using uv radiation, when microorganisms were deactivated from the surface while maintaining material and chemical parameters of individual samples. sterilization methods degrading the original properties of the material was not used. samples were inoculated with a consortium of molds and then placed at 25 ± 3 °c and monitored for 41 days. the consortium of fungi contained microorganisms from the czech collection of microorganisms: paecilomyces variotti, chaetomyum globosum, penicillium funiculosum, aspergillus terreus, penicillium brevicompactum. the evaluation of the resistance of the material took place visually on the 14th, 22nd, 29th and 41st day. sample 1 has isolated colonies of molds after 14 days of experiment. on day 22, coverage was already 80% and more (figure 3). sample 2 is very resistant to mold. the samples showed isolated mold hyphae, but they grew more out of the medium. at the end of the experiment, fungi were visible at the bottom of the flanks and isolated colonies of mold at the top (figure 4). mold colonies were already visible on day 14 on sample 3 and on day 22 the samples were completely covered (figure 5). sample 4 was the most resistant to the occurrence of mold, after 41 days only isolated colonies of mold were visible on the surface, the sides of the samples were only partially covered (approximately 30%). the mixture of sample 5 had a similar result, only the coverage rate after 41 days was several percent higher (figure 6). sample 6 had almost the entire sides covered and mold colonies were already present on the surface on day 22. the coverage rate remained almost unchanged by the end of the experiment. sample 7 prevented the growth of mold until day 41, where iso95 h. sekavová, z. prošek, p. ryparová acta polytechnica ctu proceedings figure 3. sample 1 after 14 (left) and 22 (right) days. figure 4. sample 2 after 25 (left) and 41 (right) days. figure 5. sample 3 after 11 (left) and 25 (right) days. figure 6. sample 4 (left) and sample 5 (right) after 41 days. 96 vol. 30/2021 detection of the presence of microorganisms in plasterboard figure 7. sample 6 (left) and sample 7 (right) after 41 days. figure 8. sample 8 (left) and sample 9 (right) after 11 days. lated colonies were visible, both on the surface and on the sides (figure 7). paper samples were most prone to mold growth and growth began as early as the first week (figure 8). 4. conclusions from the above results it is clear that the most risky part of gypsum plasterboards is the paper part. the best results achieve the samples of hardened gypsum slurry type h (impregnate) and hardened gypsum slurry prepared in laboratory from stucco and water. the experiment was designed to simulate a very demanding environment and in the case of the use of these materials in real constructions, the time of microbial contamination depends on the conditions, which are humidity, temperature and possible pollution as a source of nutrients for microorganisms. these results are the basis for subsequent modifications to prevent the development of microorganisms in gypsum plasterboard systems. acknowledgements this paper was financially supported by the ta cr research a project under the number tj02000235 and by czech technical university in prague under no. sgs project sgs19/148/ohk1/3t/11. the authors also thank knauf, praha ltd. for the supplied samples and materials and the center for nanotechnology in civil engineering at the faculty of civil engineering, czech technical university in prague. references [1] m. nyč. sádrokarton. grada publishing as, 2005. [2] j. herrmann. možnosti recyklace výrobkŭ na bázi sádry 2019. [3] h. sekavová, j. herrmann, z. prošek, et al. the influence of the use of recycled gypsum on the properties of gypsum products in labor conditions. acta polytechnica ctu proceedings 26:81–85, 2020. [4] r. wasserbauer. biologické znehodnocení staveb [biological damage of structures]. abf arch: praha, prague 2000. [5] n. antošová. impact of biocorrosion on the durability of etics and empirical findings about the periodocity of maintenance. slovak journal of civil engineering 21(2):21–28, 2013. [6] v. kučeriková, m. kraus, d. kubečková. analysis of the degradation of external plasters in the buildings with etics. in advanced materials research, vol. 1041, pp. 15–18. trans tech publ, 2014. [7] m. havrlik, p. ryparová. protection of wooden materials against biological attack by using nanotechnology. acta polytechnica 55(2):101–108, 2015. [8] čsn en 520+a1. sádrokartonové desky definice, požadavky a zkušební metody. praha: úřad pro technickou normalizaci, metrologii a státní zkušebnictví 2010. [9] čsn en iso 846. plasty hodnocení působení mikroorganismů. praha: úřad pro technickou normalizaci, metrologii a státní zkušebnictví 2019. 97 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.37.0010 acta polytechnica ctu proceedings 37:10–15, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the characterization of the collimated beams of fast neutrons with the clid detection system martin ansorgea, b, ∗, jan novákb, mitja majerleb, ján kozica, b a czech technical university in prague, faculty of nuclear sciences and physical engineering, v holešovičkách 747/2, 180 00 prague 8, czech republic b nuclear physics institute of the czech academy of sciences, 250 68 řež, czech republic ∗ corresponding author: ansorge@ujf.cas.cz abstract. a new detection device for the measurements of light ions (p, d, t, α) emitted as the products of the nuclear reactions induced by fast neutrons (5-33 mev) was recently developed at the nuclear physics institute of the czech academy of sciences. the main objective of the chamber-forlight-ion-detection (clid) is to produce new differential nuclear data of high interest for the material applications related to fusion and aerospace technologies and to potentially test and validate models of nuclear reactions. hereby the experimental set-up for the measurements with the clid is described in detail. the experimental characterization of the collimated fast neutron beams produced by the cyclotron-driven converter (p(35 mev)+be(2.5 mm)) is presented. in particular, the implementation of the proton-recoil-telescope technique used for neutron energy spectra determination with the clid is described. keywords: npi, collimated fast neutron beams, cyclotron u-120m, clid, proton-recoil-telescope. 1. introduction a new 1.6 m long collimator coupled to a new neutron converter (based on the p+9be reaction) driven by the isochronous cyclotron u-120m was constructed at the nuclear physics institute of the czech academy of sciences (npi-cas). collimated beams of fast neutrons are bringing new experimental possibilities, especially for on-beam measurements with semiconductor detectors which need to be shielded from intense neutron fields. the new vacuum chamber with the telescopes composed of silicon detectors allows us to perform precise measurements of double-differential cross sections of (n,cp charged particle) reactions induced by fast neutrons with kinetic energies from 5 to 33 mev. motivated by the development of future fission and fusion energy projects, the nuclear data measurements of cross sections for nuclear reactions induced by fast neutrons are of high interest. in recent years, the clid detection system was commissioned and it delivered its first experimental data. this paper describes the characterization of the collimated neutron fields available at the u-120m facility at the npi-cas. the neutron fluxes and energetic spectra were measured in the irradiation position of the clid device using the experimental set-ups with the vacuum chamber as well as with the scintillation probes. the development of the clid system follows the previous development of similar devices used at several nuclear research facilities worldwide. during past decades, the ddcs data for (n,cp) reactions induced by fast neutrons (10-200 mev) on several basic elements (c, li, o, al, si, fe, co, bi, u, etc.) were measured and published at lansce in usa [1], at cyclon in belgium [2–4], at crocker nuclear laboratory in the usa [5–7] or at svedberg laboratory in sweden where the medley chamber was constructed and used for several campaigns of ddcs measurements [8–11]. recently, the medley was moved for the new set of experimental campaigns to the ganil laboratory in france where the new linac accelerator is used for neutron production mainly in the energy region from 1 to 40 mev [12]. in 2022, new data on α-particle production double differential cross-sections on carbon were published by the chinese group from the csns facility for the energy range from 6 to 102 mev [13]. 2. general set-up for clid experiments the initial proton bunches are accelerated by the cyclotron u-120m to the kinetic energy of 34.8±0.5 mev. pulsed beams of primary protons are delivered through few meters of ion beam-line to a thin beryllium-based neutron converter where fast neutrons are produced. the neutron converter is placed very close to the 1.6 m long circular-shaped collimator (3 cm in diameter). behind the collimator, the evacuated reaction chamber is placed at an approximate distance of 2.5 m (from the neutron converter to the irradiated sample). the representative schematics of the experimental set-up is shown in figure 1. 2.1. cyclotron-driven neutron source the new neutron converter consists of 2.5 mm thick beryllium slab. the neutrons with quasi10 https://doi.org/10.14311/app.2022.37.0010 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 37/2022 clid detection system figure 1. the top-view schematics of the experimental set-up used with the reaction chamber. monoenergetic (qm) spectrum are produced by 9be(p,n) reactions. the beryllium is backed by 10 mm thick slab of graphite which is used to stop the protons. the beryllium/graphite slabs are cooled by the flow of de-mineralized water. the primary proton beam current is measured by the integration of the charge deposited by protons in the beryllium/graphite slabs. historically, the qm neutron spectra produced by beryllium targets were already studied at the npicas. a 0.5 mm thick beryllium target was bombarded by ≈30 mev protons and the neutron energy spectra were measured by the time-of-flight (tof) technique using the ne-213 scintillator [14]. two monoenergetic peaks separated by 2-3 mev were recorded on the neutron continuum. based on these results, five times thicker beryllium slab was chosen to produce only one broader peak of monoenergetic neutrons on continuous neutron background. the target station with the neutron converter is placed close to the collimator wall as shown in figure 1. a hollow steel cylinder was implemented in this wall. to shape the beam of fast neutrons, several hollow cylinders made of polyethylene and steel pieces were inserted in the inner part of the collimator cylinder. the rest of the collimator wall consists of the mixture of iron and water. 2.2. clid device inspired by the medley setup [8], a new vacuum chamber for the detection of light ions (p, d, t, 3he, α) produced in nuclear reactions induced by fast neutrons (up to 33 mev) was constructed at the npicas. the chamber allows the measurements of doubledifferential cross sections for reactions (n, cp) with respect to the angle and incident energy of the interaction. the time-of-flight method is used for the energy determination of the primary neutron and remotely controlled rotating table is used for angle differentiation. the central sample holder can be loaded with up to four different targets which can be switched during the on-beam measurement remotely. using the clid, the neutron flux and energy spectrum were measured by the proton-recoil-telescope (prt) method. the polyethylene foil (pe,ch2 250 µm thick and 2 cm diameter) mounted on the strip of 2.5 µm thin mylar foil was placed at the center of the chamber and irradiated. the elastically scattered protons were recorded using the telescopes. a carbon sample of the same diameter and half of the thickness was measured in the same position in order figure 2. the photography of the internal set-up of the clid placed behind the collimator. 11 m. ansorge, j. novák, m. majerle, j. kozic acta polytechnica ctu proceedings to subtract the background proton events from (n,xp) reactions on carbon in pe. at last, an yttrium sample was irradiated and analyzed using the activation technique. the four si-based telescopes were placed on the rotating table around the irradiated sample, see the figure 2. each telescope consisted of two planar si detectors, the thin one to record the energy loss (de) of the passing charged particle and the thick one to absorb the rest of the particle energy (e). in the case of the prt measurement with the pe foil, the telescope consisted of the 150 µm and 5 mm thick detectors. the solid angle between the pe sample and the telescope was restricted by the circularly shaped aperture with the inner diameter of 8 mm. this solid angle was calculated by numerical integration technique based on the method presented by s. pommé [15] and the resulting absolute value of 0.00255 sr was obtained. the telescope was placed at the laboratory angle of 45° with respect to the neutron beam axis. the data acquisition chain consisted of the caen preamplifiers and v1730 digitizer. the power supply modules delivering the highand low-voltage were manufactured by caen as well. 3. prt measurement of the neutron spectrum during the approximately 2-hour-long irradiation of the ch2 target at the beam current of approx. 20 µa the light charged particles were recorded by the prt at the laboratory angle of 45° and proton events were discriminated by the de-e method which is graphically represented by the 2-d plot in figure 3. the energy calibration of the si-detectors has been performed using the radionuclide α-sources with kinetic energies of α-particles up to 8 mev. figure 3. particle type separation for prt data. the background runs were measured with the carbon and empty-holder targets. the background and natc(n,p) proton events were correctly subtracted and the proton spectrum from h(n,p) elastic scattering was reconstructed, see the blue data-set in figure 4. to obtain the energy spectrum of protons produced in the pe sample, the unfolding procedure (based on roounfold package [16]) was implemented to correct the measured spectra for energy spreading and shifting figure 4. proton energy spectra measured by prt (blue) and corrected proton energy spectra (red). due to the energy loss in the pe sample and due to the variations of the recoil angle within the solid angle of the detector. the response function used for the unfolding process was generated using the montecarlo code geant4 [17] and the previously created clid model [18]. the simulated response matrix is plotted in figure 5 and represents the response of the prt set-up of the clid to the theoretical irradiation by the mono-energetic neutrons which are elastically scattered on hydrogen nuclei in pe foil. figure 5. geant4 simulated response function for the set-up with ch2 target and prt. the kinetic energy of protons is recalculated into the energy of interacting neutrons based on the welldefined kinematics of elastic scattering. the following equation is used to calculate the single neutron spectral point in each energy bin of the final neutron spectrum: nn = np ωlabσdif f habwc [n/cm2/mev/c], (1) • np – number of protons detected in the energy bin, • ωlab – solid angle (in lab system) from ch2 target to telescope, • σdif f = ∂σ∂ωlab ∂en – endf/b-vi double differential cross-section for h(n,p), • ha – number of hydrogen atoms in the ch2 sample, • bw – the width of energy bin for which the spectral point is calculated, • c – the charge of primary proton beam integrated over the period of measurement. 12 vol. 37/2022 clid detection system 4. activation measurement two 12-hours-long irradiations of the yttrium foil were performed. yttrium is commonly used as activation detector material and has a well-known (n,xn) reaction cross sections. the parameters for two relevant neutron-induced reactions are listed in table 1. the reaction rates for the production of the 88y isotope were successfully measured (with hpge), while the production of 89y(n,3n)87y could not be measured due to the long cooling times of the yttrium foil compared to the relatively short half-life period. reaction q-value[mev] threshold [mev] half life [days] 89y(n,2n)88y -11.48 11.61 106.65 89y(n,3n)87y -20.83 21.07 3.33 table 1. basic reaction parameters relevant for activation technique assessment. the overview of experimental and evaluated crosssection data for the 89y(n,2n)88y reaction plotted by janis [19] is shown in figure 6. the measured figure 6. experimental and evaluated cross-section data for 89y(n,2n)88y reaction plotted by janis [19]. reaction rates can be used to control the correctness of the absolute values in the neutron spectra measured by the prt technique as will be shown in the following section 5. 5. results the neutron energy spectrum at the irradiation position of the clid was measured by the prt technique and calculated using the eq. 1 together with the measured proton data presented in figure 4. the resulting neutron energy spectrum is plotted in figure 7. the reaction rates (number of produced nuclei per 1 g of the sample and per 1 primary particle) for the production of 88y isotope were measured by the hpge detectors. besides the activation technique described in the previous section 4, the reaction rate was also calculated by the integration of the known cross-section σ(e) (figure 6) with the prt-measured figure 7. spectral neutron flux measured by the prt method with the clid. neutron flux ϕ(e) over the energy range of the neutron energy spectrum in figure 7: rr = ∫ emax emin σ(e)ϕ(e)de, (2) where e is the neutron kinetic energy. the neutron cross-section data for σ(e) were taken from the eaf2010 [20] data library. the absolute values for the rr quantity calculated by both approaches (activation method, prt method) are listed in the table 2. the resulting normalization factor is listed as well and the disagreement between the two measured values is further explained in the conclusion of this article. rr by activation method rr for prt method normalization factor [/g/c] [/g/c] [-] 6.71e+08 4.85e+08 1.3844 table 2. the reaction rates calculated for 89y(n,2n)88y reaction and their ratio. in the figure 8 the final neutron energy spectrum measured by the prt technique and corrected by the normalization factor deduced from activation measurement with yttrium sample is shown. an independent measurement with the ne213 scintillator in the tof mode was performed to obtain the neutron spectral flux at the irradiation position of the clid. the methodology of the tof measurement is beyond the scope of this paper [21, 22]. the resulting tof neutron spectrum is shown in figure 8 for comparison. furthermore, the mcnpx [23] calculation was performed to simulate the neutron flux in the clid irradiation position. the complete geometry with the collimator was used in the mcnpx simulation, the beryllium target backed with the carbon slab and the layer of coolant in the target station was modelled. the neutron production was simulated for the p(34.8 mev)+9be(2.5 mm thick) reaction using the jendl4.0/he data library [24, 25]. comparison between experimentally measured neutron fluxes and mcnpx simulation is shown in figure 8 as well. 13 m. ansorge, j. novák, m. majerle, j. kozic acta polytechnica ctu proceedings figure 8. spectral neutron fluxes measured by prt method and normalized by 88y production rates (red), measured by the ne213 (green) and simulated by the mcnpx code using the jendl4.0/he data library (blue)[24, 25]. for all three energy spectra plotted in figure 8, the absolute values of the qm-peak neutrons measured or simulated at the clid irradiation position were integrated, see the table 3. the lower integration limits were 20 mev, 21 mev and 23 mev for data sets from ne213, prt and mcnpx respectively. both measurements (ne213, normalized prt) are in very good agreement with each other. on the other hand, the mcnpx simulation gives roughly 30 % lower values for the qm-peak integral compared to the measurements. this disagreement with the mcnpx results is probably due to the poor precision of the implemented cross-section data for the production of neutrons in the jendl4.0/he library. the jendl4.0/he has however better description of the neutron production than the other existing library endf7prot [26] and therefore was chosen for this comparison. ne213 prt mcnpx tof normalised simulation peak neut. [n/cm2/c] 1.10e+11 1.11e+11 8.067e+10 table 3. integral values of qm neutrons at the clid irradiation position. 6. conclusions this paper describes the overall experimental set-up with the new vacuum chamber. the implementation of the prt technique for neutron flux determination with the clid was described in detail. during the one weekend-long on-beam experiment with the clid, the ch2, 89y and natc targets were irradiated and the measured data were used for the characterization of the fast neutron spectrum in the collimated beam. the detailed analysis of the experimental data has shown that some events were randomly lost during the data acquisition or their energy deposition values were falsely recorded. the evidence for this effect can be found around the proton band in figure 3. this issue was already addressed and solved for future experiments. the experimental prt data presented in this paper were however compromised and consequently had to be normalised. the normalisation factor was deduced from the activation measurement of the irradiated yttrium sample. the 89y(n,2n)88y reaction has very well known cross section (figure 6) in the energy range from 11 mev to 36 mev. the correctness of the activation measurement was confirmed by the tof measurements performed with the organic scintillator ne213. the rr = 6.711e08 88y/g/c value calculated for the neutron energy spectrum (green in figure 8) measured by the ne213 matches perfectly with the rr value presented for activation in table 2. the tof spectrum agrees with the normalized prt neutron energy spectrum. the integral values for qm-peak neutron fluxes (as presented in table 3) are of the key importance for every experiment on collimated fast neutron beams produced on 2.5 mm thin beryllium target by the cyclotron u-120m at energies around 34.8 mev. the overall relative uncertainty of the measured data in table 3 is expected to be below 10 %. the accuracy of the mcnpx/jendl4.0/he predicted values is lower because of the worse implementation of the neutron production in the jendl4.0/he evaluation. the broad qm peak from the p+be reaction on 2.5 mm thin beryllium target is composed of two narrow peaks which originate from the ground and excited states of the remaining 9b nucleus [14]. the absolute values of cross section and the intensity ratio between these two inner peaks is not implemented correctly into the evaluated data libraries. the production of neutrons is however better described in jendl4.0/he than in the other existing libraries generally used with mcnpx. acknowledgements this project has received funding from the european union’s horizon 2020 research and innovation programme under grant agreement no 847552. this work was supported by op rde, meys, cz under the project canam, ef16_013/0001812. computational resources were supplied by the project "einfrastruktura cz" (e-infra cz lm2018140) supported by the ministry of education, youth and sports of the czech republic. this work was supported by the grant agency of the czech technical university in prague, grant no. sgs18/145/ohk4/2t/14. references [1] p. lisowski, c. bowman, g. russell, s. wender. the los alamos national laboratory spallation neutron sources. nuclear science and engineering 106(2):208– 218, 1990. https://doi.org/10.13182/nse90-a27471. 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https://doi.org/10.11484/jaea-conf-2016-004 acta polytechnica ctu proceedings 37:1–6, 2022 1 introduction 2 general set-up for clid experiments 2.1 cyclotron-driven neutron source 2.2 clid device 3 prt measurement of the neutron spectrum 4 activation measurement 5 results 6 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0363 acta polytechnica ctu proceedings 33:363–369, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague properties of concrete using crushed stone powder with various specific surface areas takeju matsukaa, ∗, mariko suzukib, haruhiko matsushitac, hiroshi yokotad, koji sakaie a national institute of technology, kumamoto college, 2627 hirayama-shinmachi, yatsushiro, kumamoto, 866-8501, japan b kobe university, 1-1 rokkodai-cho, nada, kobe, 657-8501, japan c chuosaiseki co., ltd., 856-4 hara, takatsuki, 569-1051, japan d hokkaido university, kita 13, nishi 8, kita-ku, sapporo, hokkaido, 060-8628, japan e japan sustainability institute, sapporo, hokkaido, 065-0021, japan ∗ corresponding author: matsuka@kumamoto-nct.ac.jp abstract. many countries use crushed stone and crushed sand as the aggregate for concrete. however, a large amount of crushed stone powder is produced as by-products in the manufacturing process. unless effective utilization of crushed stone powder is considered, it is not a sufficient measure in terms of sustainability. in this study, the properties of concrete using crushed stone powders with different specific surface areas are investigated. as the results, a strong correlation was found between the unit water content of concrete and the total surface area of the crushed stone powder contained in the concrete. furthermore, it was concluded that the correlation between the strength index, the co2 index, and the virgin resource index of concrete using crushed stone powder depends on the types of crushed stone powder and crushed sand. keywords: crushed stone powder, fine aggregate, fresh concrete, hardened concrete, specific surface area. 1. introduction crushed stone and crushed sand are often used as the aggregate for concrete in many countries. in the future, concrete aggregates would rely on crushed stone and crushed sand because natural aggregate, such as sea sand is probably prohibited from collecting. the production methods of crushed stone and crushed sand are classified into two types, such as a dry process type and a wet process type. in the dry process type, a large quantity of crushed stone powder is simultaneously produced as by-products. the japan crushed stone association estimates that 12 million tons of crushed stone powder are produced annually. in the wet process type, although crushed stone powder and clay fraction can be removed, sludge water and a dehydrated cake must be disposed. in recent years, the dry process type has become more popular because of its production efficiency and ease of handing waste disposal [1]. however, unless effective utilization of crushed stone powder is considered, a part of the natural resources will be unused, which is not a sufficient measure from the perspective of sustainability. in the world, studies on the effective use of crushed stone powder in concrete have been carried out, but they are few [2, 3]. in 2009, the quality requirements of crushed stone powder were standardized by the japanese industrial standards, jis a 5041. however, no specification is available on the specific surface area of the crushed stone powder. the authors studied the fluidity of mortar using crushed stone powders with various specific surface areas [4]. they found that the watercement ratio of mortar has a high influence on the fluidity of mortar with the crushed stone powder and, if the crushed stone powder with a large specific surface area is used, the fluidity of the mortar decreases. however, these findings are limited to mortar. the properties of concrete using crushed stone powders are not clear. in this study, the properties of the fresh and hardened concrete using crushed stone powders with various specific surface areas were examined. furthermore, co2 emission and the amount of virgin resource input for each concrete mix were estimated. 2. manufacturing of crushed sands and crushed stone powders figure 1 shows the simplified flow chart of the production processes of crushed sand and crushed stone powder in this study. in the dry process type, crushed stone powder type a and crushed sand ds are produced at first. then crushed stone powder type a is sorted by specific surface area as type b and type c. ds contains crushed stone powder type d. ds150 is produced by removing crushed stone powder of 150 µm or less contained in ds. in the wet process 363 https://doi.org/10.14311/app.2022.33.0363 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en t. matsuka, m. suzuki, h. matsushita et al. acta polytechnica ctu proceedings figure 1. production processes of crushed sands and crushed stone powders. figure 2. grain size accumulation curve of crushed sand. type, crushed sand ws and a dehydrated cake are produced. the dehydrated cake is, however, not the target of this study. ws contains crushed stone powder type e. ws150 is produced by removing crushed stone powder of 150 µm or less contained in ws. finally, the crushed sand cs is produced by mixing ws and crushed stone powder type b. in this study, the effect of the combination of crushed stone powder and crushed sand on the properties of concrete is examined. table 1 gives the types and physical properties of crushed sand and crushed stone powder. 3. outline of experiment 3.1. materials table 1 gives the types and physical properties of concrete materials used in this study. as mentioned earlier, in this experiment, two types of dry type crushed sand, three types of wet type crushed sand, and five types of crushed stone powder were used. the specific surface area of crushed stone powder was measured using a laser diffraction particle size analyzer. the specific surface area of the crushed stone powder was figure 3. grain size accumulation curve of crushed stone powder . in the range from 6360-22337 cm2/g. figures 2 and 3 show the grain size accumulation curves of crushed sand and crushed stone powder, respectively. the particle size of crushed stone powder is less than 150 µm as per jis a 5041. 3.2. production and mix proportions of concrete a total of 12 mixtures were planned to produce. concrete was produced in a laboratory at a temperature of 20 řc with a 60 l twin-shaft batch mixer. cement, fine aggregate, crushed stone powder, and coarse aggregate were first dry mixed for 15 seconds, and then further mixed for 90 seconds after adding water and chemical admixture. the water-cement ratio (w/c) was 60% for all the mixtures. the amount of air entraining (ae) water reducing agent was 1.2% to cement mass. table 2 lists the mix proportions of the concrete. when mixing concrete, the slump target was 12.0 ± 2.0 cm and air content target was 4.5 ± 1.0%; the slump and air content were adjusted based on the unit quantities of water and ae agent, respectively. 364 vol. 33/2022 crushed stone powder concrete materials type symbol physical properties water tap water w density: 1.00 g/cm3 cement ordinaryportland cement c density: 3.16 g/cm3, specific surface area: 6315 cm2/g (3300 cm2/cm3) arithmetic average particle diameter: 19.4 µm fine aggregate crushed sand (dry type) ds density: 2.62 g/cm3, absorption: 1.80%, finess modulus: 2.85 content of 75 µm or less: 4.0% content of 150 µm or less: 8.0% ds150 density: 2.66 g/cm3, absorption: 1.48%, finess modulus: 2.97 content of 75 µm or less: 0% content of 150 µm or less: 0% crushed sand (wet type) ws density: 2.65 g/cm3, absorption: 1.76%, finess modulus: 2.84 content of 75 µm or less: 2.3% content of 150 µm or less: 7.8% ws150 density: 2.66 g/cm3, absorption: 1.42%, finess modulus: 2.99 content of 75 µm or less: 0% content of 150 µm or less: 0% cs density: 2.63 g/cm3, absorption: 1.79%, finess modulus: 2.80 content of 75 µm or less: 5.0% content of 150 µm or less: 9.1% crushed stone powder type a density: 2.684 g/cm3 specific surface area: 7120 cm2/g arithmetic average particle diameter: 35.2 µm type b density: 2.682 g/cm3 specific surface area: 6460 cm2/g arithmetic average particle diameter: 31.9 µm type c density: 2.687 g/cm3 specific surface area: 22337 cm2/g arithmetic average particle diameter: 3.6 µm type d density: 2.783 g/cm3 specific surface area: 16825 cm2/g arithmetic average particle diameter: 17.2 µm type e density: 2.772 g/cm3 specific surface area: 12914 cm2/g arithmetic average particle diameter: 25.8 µm coarse aggregate sand stone g1 maximum size: 20 mm, density: 2.70 g/cm 3 absorption: 0.47%, finess modulus: 7.08 g2 maximum size: 15 mm, density: 2.69 g/cm 3 absorption: 0.76%, finess modulus: 6.35 chemical admixture ae water reducing agent − complex of lignin-sulfonic-acid compound and polycarboxylic acid ether ae agent − denatured rosin acid compound-system anionicsurface active agent table 1. constituent materials of concrete. 365 t. matsuka, m. suzuki, h. matsushita et al. acta polytechnica ctu proceedings quantity of material per unit volume of [kg/m3] no. w/c w c ds150 ws ds 150 ws cs type b type c g1 g2 ae water reducing agent ae agent [%] ! kg/m3 " ! kg/m3 " 1 60 162 270 872 410 613 c×1.2% 0 2 162 270 872 0 3 165 275 846 0.275 4 168 280 844 0.280 5 165 275 850 0.138 6 165 275 830 26 0.413 7 170 283 777 51 0.992 8 165 275 805 52 0.668 9 168 280 787 51 0.980 10 173 288 766 50 1.009 11 170 283 786 51 0.992 12 180 300 668 118 1.500 table 2. mix proportions of concrete. figure 4. amount of crushed stone powder of 150 µm or less for each mix proportion. figure 4 shows the amount of crushed stone powder of 150 µm or less for each mix of concrete. the amounts of crushed stone powder for each mix proportion were in the range of 0 − 170.5 kg/m3. they are included the amount of crushed stone powder in the crushed sand as well. figure 5 shows the unit water content of each mix proportion. in addition, this figure shows the crushed stone powder ratios of 150 µm or less and 75 µm or less. the crushed stone powder ratio is defined in this study as the ratio of the absolutely dry mass of crushed stone powder mass to that of the fine aggregate. 3.3. test methods for properties of concrete for fresh concrete, slump, air content, and bleeding were measured as per jis a 1101, a 1128, and a 1123, respectively. for hardened concrete, compressive strength and youngs modulus were measured as per jis a 1108 and a 1149, respectively. the shape figure 5. unit water content of each mix proportion. of the specimens used for the test of hardened concrete is a cylinder with a diameter of 100mm and a height of 200mm. the number of samples in the test at each age is three. 4. results and discussions 4.1. unit water content figure 6 shows the relationship between the crushed stone powder ratios of 150 µm or less and 75 µm or less and the unit water contents. a correlation exists between the crushed stone powder ratios and the unit water content. the unit water content is specified as 175 kg/m3 or less in the standard specifications for concrete structures (japan society of civil engineers (jsce)). the fine particle content of 75 µm or less is specified as 9% or less in jis a 5005. these results indicate that the crushed stone powder ratio of 75 µm or less should be less than 15% to satisfy the jsce specifications. therefore, it has been suggested 366 vol. 33/2022 crushed stone powder concrete figure 6. relationship between crushed stone powder ratio and unit quantity of water. figure 7. relationship between surface area of crushed stone powder and unit quantity of water. that the specification of jis a 5005 (9%) of the fine particles content can be moderated for concrete. in this study, the total surface area of the crushed stone powder with particle sizes of 150 µm or less was calculated, and the relationship between the total surface areas of the crushed stone powder and the unit water content is shown in figure 7. it was found that the total surface areas of crushed stone powder have a stronger relationship than the crushed stone powder ratio because the determination coefficient is improved. therefore, it can be concluded that the specific surface area of the crushed stone powder is an effective index for evaluating the fluidity of concrete. 4.2. amount of ae agent figure 8 shows the relationship between the amount of ae agent and the crushed stone powder ratio. the amount of ae agent increases with increase in the crushed stone powder ratio. it is well known that air is difficult to be entrained when concrete has fine particles of 0.15 mm or less. the results of this study confirms it. figure 8. amount of ae agent. figure 9. amount of bleeding. 4.3. amount of bleeding figure 9 shows the relationship between the amount of bleeding and the crushed stone powder ratio. to ensure the water-tightness of concrete, the amount of bleeding is specified as 0.3 cm3/cm2 or less by the japanese architectural standard specification 5 (jass 5) [5]. the experimental results indicate that the amount of bleeding is 0.3 cm3/cm2 when the crushed stone powder is 8.5%. based on this result, controlling the amount of crushed stone powder is required to reduce bleeding. 4.4. compressive strength and young’s modulus figure 10 shows the relationship between the crushed stone powder ratio and the compressive strength of concrete. this figure indicates that the compressive strength decreases with increase in the crushed stone power ratio. when the crushed stone powder is used, the compressive strength decreases by up to 21.4% and by 20.0% at 7 days and 28 days, respectively. figure 11 shows the relationship between the compressive strength and the young’s modulus of concrete. this figure also shows the calculated results using the equation specified in the jsce standard 367 t. matsuka, m. suzuki, h. matsushita et al. acta polytechnica ctu proceedings figure 10. compressive strength. specifications for concrete structures [6]. the ratio of the experimental value to the calculated value of the young’s modulus of concrete were in the range 0.99 to 1.15. it can be said that the calculated results estimate the experimental results well almost appropriately. 5. environmental assessment 5.1. outline the importance of sustainability has recently been emphasized in the concrete industry [7]. sustainability is considered from the following three aspects: social, economic, and environmental. in the future, all technologies in the concrete industry will be evaluated from the perspective of sustainability; thus, all technologies in the concrete industry must be advanced based on the above-mentioned three sustainability aspects. in particular, reducing co2 emissions is currently an urgent issue in the world. in this study, the effect of using the crushed stone powder is evaluated from the perspective of sustainability, such as safety in terms of compressive strength and environmental performance in terms of co2 emissions and virgin resource input. 5.2. calculation methods the amounts of co2 emissions and virgin resource input for each mix proportion were calculated. the inventory data relating to the co2 emissions are obtained from "environment text of concrete (draft)" [8]. table 3 gives the unit-based co2 emissions in this study. this study additionally considered co2 emissions associated with kerosene consumed in the mixing and classification of crushed stone powder. however, the co2 emissions associated with electricity was not considered, because the breakdown of the quantity of electricity has not been clarified. the amount of virgin resource input was determined by multiplying the virgin resource input ratio of each material by a unit amount per 1 m3 of concrete. the virgin resource input ratio of cement figure 11. relationship between young’s modulus and compressive strength. items unit-based co2 emissions water 0.348 kg-co2/m3 cement 769 kg-co2/t crushed sand 3.9 kg-co2/t kerosene 2.50 kg-co2/l table 3. co2 emission unit. is 0.811 [9], and that of crushed stone powder a, b, and c is 0 because they are by-products that can be discarded. the virgin resource input ratios of the other materials are set at 1.0. 5.3. amounts of co2 emissions and virgin resource input figure 12 and 13 show the amounts co2 emissions and the virgin resource input, respectively. the co2 emissions increase with increase in the crushed stone powder, because kerosene is consumed to produce the crushed stone powder. in contrast, the amount of virgin resource input decreases with increase in the crushed stone powder. the amount of co2 emissions and the virgin resource input were in the range of 215 − 244 kg-co2/m3 and 2115 − 2276 kg/m3, respectively. 5.4. sustainability evaluation figure 14 shows an example of the correlation between the strength index, co2 index, and virgin resource index. each index was defined based on the respective values of the reference concrete (no.1 mix proportion). the larger the area of the triangle, the higher the overall performance. the results indicate that the relationships between the strength index, co2 index, and virgin resource index of concrete using the crushed stone powder depend on the types of crushed stone powder and crushed sand. it is concluded that evaluation from the perspective of environmental sustainability is also important for the mix proportion design of concrete using crushed stone powder. in addition, further data 368 vol. 33/2022 crushed stone powder concrete figure 12. co2 emissions. figure 13. amount of virgin resource input. accumulation is needed to generalize the various performances of the combinations of the specific surface area of crushed stone powder and the type of crushed sand. 6. conclusions the following conclusions were drawn from this study: • the specific surface area of crushed stone powder greatly affects the fluidity of concrete. • increasing the mixing ratio and the surface area of the crushed stone increase the unit water content of the concrete. however, controlling the amount of crushed stone powder is necessary for reducing the bleeding of concrete. • using crushed stone powder decreases the compressive strength of concrete up to 21.4 • the crushed stone powder reduces, the virgin resource input, but increases the co2 emissions. • the relationships between the strength index, co2 index, and virgin resource index of concrete using crushed stone powder depend on the type of crushed stone powder and crushed sand. figure 14. sustainability evaluation. references [1] t. mano, y. tsuji, f. tomosawa, et al. properties of fresh and hardened concrete using crushed stone powder as part of fine aggregate. concrete journal 46(11):18-24, 2008. https://doi.org/10.3151/coj1975.46.11_18. [2] t. h. song, s. h. lee, b. kim. recycling of crushed stone powder as a partial replacement for silica powder in extruded cement panels. construction and building materials 52:105-15, 2014. https: //doi.org/10.1016/j.conbuildmat.2013.10.060. [3] s. p. s. rajput. an experimental study on crushed stone dust as fine aggregate in cement concrete. materials today: proceedings 5(9):17540-7, 2018. https://doi.org/10.1016/j.matpr.2018.06.070. [4] suzuki m, matsuka t, yokoyama m and matsushita h 2019 influence of crushed stone powder with different specific surface area on fluidity of crushed sand mortar irrigation, drainage and rural engineering journal 87 pp 43-48. [5] architectural institute of japan. japanese architectural standard specification jass 5 reinforced concrete work, 2018. [6] japan society of civil engineers. standard specification for concrete structures -design-, 2018. [7] k. sakai, h. yokota.sustainability design of concrete structures gihodo shuppan co., ltd, 2016. [8] japan concrete institute. environment text of concrete (draft), 2016. [9] h. saito, k. sakai, y. suzuki, et al. proposal of evaluation method of low carbon concrete using fly ash and ground granulated blast-furnace slag. proceedings of the japan concrete institute 36:1852-1857, 2014. 369 https://doi.org/10.3151/coj1975.46.11_18 https://doi.org/10.1016/j.conbuildmat.2013.10.060 https://doi.org/10.1016/j.matpr.2018.06.070 https://doi.org/10.14311/app.2022.33.0045 acta polytechnica ctu proceedings 33:45–51, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague sustainability of concrete pavements considering traffic and de-icing agents rolf breitenbücher, robin przondziono∗ ruhr university bochum, department of civil and environmental engineering, institute for building materials, universitätsstrasse 150, 44801 bochum, germany ∗ corresponding author: robin.przondziono@ruhr-uni-bochum.de abstract. in concrete pavements, special conditions, that increase the likeliness of damaging reactions, as for example an alkali-silica reaction (asr), prevail. especially to the superposition of microstructural degradation caused by cyclic loading with an external alkali supply can influence the sustainability negatively. concrete pavements are subjected to cyclic loadings by traffic and climate changes. these cause microcracks (about 5 µm) within the concrete matrix during service. additionally, an asr in pavements is promoted by externally supplied alkalis (de-icing agents). by superposition of both effects the alkali capacity close to the asr-reactive aggregate aggregate is increased substantially as the externally supplied alkalis can easily penetrate through the microcracks. thus, both effects intensify the asr in concrete pavements. within cooperative research projects the different interdependent influencing factors for a damaging asr in concrete pavements are studied by experiments as well as by numeric modelling. on the micro-level the asr-related processes within the aggregate, such as gel-formation or ion-transport, are investigated. on the meso-level, the project focuses on the characterization of degradation effects in the concrete microstructure due to cyclic loading. further, special attention is paid to the transport behavior of fluids in such pre-damaged concrete structures. a significant increase of the penetration depth of alkaline solutions could not only be observed at different stages of progressing degradation. it becomes also obvious that the ingress of externally supplied alkalis is enhanced by the overrunning traffic. finally, on the macro-level, the risk of an asr-damage is assessed. keywords: degradation, durability, transport. 1. introduction concrete pavements are subject to static, cyclic and dynamic influences. in order to assess theses loads, different stress frequencies should be considered. due to changing climate conditions, thermally and hygrally-induced deformations are inevitable. if those deformations are restrained, restraint stresses are caused that depend both on the degree of restraint and on the stiffness of the structural component [1? ]. these stresses are changing when exposed to varying influences. in this context, it has to be distinguished between daily temperature changes (day night) and seasonal temperature and moisture changes (summer winter). these effects are characterized as lowfrequent. apart from these low-frequent influences, high-frequent load-dependent stresses due to traffic also act on a concrete structure. especially concrete pavements and likewise wind energy plants are exposed to a high number of load cycles day by day. these effects should be considered as high-frequent. such cyclic influences only rarely cause macrocracking, which would lead to a fatigue failure of the whole structure. however, due to these loads fine cracks might form in the concrete’s microstructure, which may then lead to the gradual degradation of the concrete properties, while, at the same time, the material’s stiffness decreases [2–4]. comparative ultrasonic runtime measurements are an appropriate non-destructive method to assess these effects [5]. the extent of degradation is dependent on two factors: the number of load cycles and the magnitude of stresses. this relation is usually described as the ratio between the stress σo and the adequate strength f. in pavements, in which degradation is caused by compressive as well as tensile stresses, the tensile properties are more relevant (σo/fct, f l). at the same time, it is essential to superpose the actual stresses with those that are already active in the concrete itself, i.e. residual stresses caused by former hydration heat release [1]. as suggested by holmen [6], at a stress-ratio of σo/fct, f l, below approx. 40 %, microcracks will not turn into macrocracks, even after millions of load cycles. at a stress-ratio above 65 %, fatigue failure of the concrete may occur considerably earlier (already after a few 10.000 load cycles). nevertheless, the formation of microcracks due to cyclic loading plays a crucial role for the durability of concrete pavements. as compared to undamaged concrete structures, such microstructural degradation alters the penetration behavior of fluids into concrete 45 https://doi.org/10.14311/app.2022.33.0045 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en rolf breitenbücher, robin przondziono acta polytechnica ctu proceedings figure 1. stresses and cracks in concrete pavements due to negative gradients of deformation [7]. [8]. in general, the transport of water in porous media depends on the size, type and form of the pores or microcracks. in addition, it also depends on the transport mechanisms and the physical laws that are effective in the respective pores / microcracks (i.e. chromatography effect). apart from diffusion, capillary transport is essential when it comes to fluid transport processes, since dissolved alkalis penetrate much deeper into the interior of the concrete structure and may hence increase the asr-damage potential. the penetration of alkalis into undamaged concrete structures and their influence on damaging asr have been sufficiently studied [9, 10]. however, the findings of these studies are not definitive. usually, an increased alkali concentration in the concrete’s nearsurface zone could be observed for specimens stored in nacl-solution. asr as a result of such an increased alkali concentration could only be observed, if the phvalue of the concrete was comparatively high [10, 11]. accordingly, the asr-risk was significantly lower in concretes that used cement which contained ground granulated blast furnace slag (ggbs), because it consumes ca(oh)2 in a latent hydraulic and pozzolanic reaction and thus lowers the ph-value [11, 12]. in this context, it should be noted that exhaustive studies on the interaction of an increased alkali ingress due to microstructural degradation of cracks with a width between 1 and 10 µm in relation to asr progression are yet to be performed; previous studies were conducted on undamaged specimens. related studies have shown, however, that the load of the surface traffic increases the penetration depth of alkalis and correspondingly, the asr-potential is promoted as well. 2. influencing factors on concrete pavements 2.1. degradation due to the continuous construction of concrete pavements, deformations in them are practically completely restrained. in the case of thermal or hygral changes, which extent over the complete crosssection, longitudinal restraint stresses are generated. during the wintertime, tensile stresses are raised while during summertime adequate compressive stresses are induced. although, to some extent a movement is enabled in transverse direction, restraint stresses cannot be excluded completely in this case. such longitudinal restraint stresses, constant over the complete cross-section, can result in throughcracks with nearly constant width. besides these, bending stresses can also be generated by a temperature gradient. in the case of a cooler surface, the slab would tend to warp. this however is prevented by its own weight, so that in the surface zone tensile stresses are provoked, while in the lower zone (next to the subbase) compressive stresses develop. if these bending stresses exceed the concrete flexure strength, uniform cracks arise with a larger crack width at the surface. furthermore, map cracking can be provoked by residual stresses (often also named as "eigenstresses"), when a non-linear temperature gradient arises. in this case, the net-like cracks extend only to a depth of a few millimetres. an overview of stresses acting on concrete pavements and corresponding potential damages is given in figure 1. microstructural degradation as a result of the concrete’s cyclic loading can be assessed on the basis of ultra-sonic runtime measurements of the longitudinal acoustic wave close to the concretes surface. based on the results of the ultrasonic runtime measurements, 46 vol. 33/2022 concrete pavements sustainability figure 2. relative dynamic e-modulus of the service lane (ss) and the driving lanes (1.fs and 2. fs) [13]. the relative dynamic e-modulus can be calculated [13]. thereby, comparative ultrasonic runtime measurements are an appropriate non-destructive method to assess these effects. sievering [13] measured the surface-near ultrasonic runtime on concrete pavements of german highways. herein the service lane was compared to the driving lanes, as it was exposed to significantly less cyclic traffic loads. correspondingly, the ultrasonic runtime was higher in the driving lanes and thus the calculated relative dynamic emodulus significantly lower in those lanes (figure 2). this indicates a higher microstructural degradation within the driving lanes, which is caused by the additional traffic loads compared to the service lane. 2.2. external alkali supply in common concrete structures an asr can be prevented by the limitation of the alkali content in the cement as this is normally the single alkali source. in concrete pavements, however, the external alkali supply by de-icing salts plays an important role on the asr-gel-formation as so, the total amount of reactive alkalis is increased significantly. in this context, also a modification in the application of the de-icing salt is relevant. in former times salt was dispersed, only when ice or snow were already present on the pavement. for some years these agents have been already applied, when such conditions were forecast. that way, the alkalis are in touch with the concrete surface in a stage at which the concrete is well absorptive. the absorption in the concrete is particularly enhanced by the above mentioned microcracks. further the de-icing agents are pressed quite intensely into the concrete microstructure by the overrolling traffic. hence, in combination with the prophylactic application of nacl the external support of alkalis into concrete pavements currently is significantly intensified in comparison with the past. 3. experimental investigation 3.1. degradation in experimental studies, large concrete beams measuring 180 × 50 × 27cm3 were exposed to cyclic stresses. all beams were produced using a typical pavement concrete mix according to the german guideline "tl beton-stb 07" [14]. as for a cement, cem i 42 n with a maximum aggregate size of 22ămm was chosen. finally, the concrete surface was broom finished. the concrete beams were exposed to cyclic stresses at a minimum age of 56 days. the tests were conducted in a four-point-flexural setup, in which the chosen upper and lower stresses represent different stress conditions of superposing loads acting on a concrete pavement. slowly recurring thermal restraint stresses were to be superposed with high-frequent cyclic traffic stresses. thus, an upper stress level and a corresponding lower stress level (σmax and σmin) could be derived from these boundary conditions [6]. the study yielded the following variations: • maximum stress to bending tensile strength (σo/fct, f l = 0.35; 0.50; 0.60; 0.70). • number of load cycles (n = 0; 1.0; 2.0; 5.0; 10.0 million) figure 3 shows the development of the relative dynamic e-modulus for the different stress-ratios as well as over the course of different load cycles as an average of three specimens each. as expected, the rel. dynamic e-modulus decreased stronger with a higher stress-ratio. after five million load cycles, the remaining relative dynamic e-modulus amounted to 47 rolf breitenbücher, robin przondziono acta polytechnica ctu proceedings figure 3. decrease of the rel. dyn. e-modulus of beams after 5 million load cycles and at different stress-ratios. rel. dynamic e-modulus [%] 100 92.7 91.5 86.5 80.2 70.2 number of cracks [-] 12 74 104 128 203 403 average crack width [µm] 6.6 5.2 5.3 4.8 5.7 4.6 average crack length [µm] 1,400 1,600 1,000 1,400 1,600 1,500 average crack cross section [µm2] 13,400 8,100 6,100 6,700 9,200 8,500 total cross section [µm2] 160,000 600,000 630,000 860,000 1,800,000 3,450,000 table 1. crack characteristics. about 91 % for a stress ratio of σo/fct, f l = 0.35 and to about 87 % and 75 % for σo/fct, f l = 0.50 and σo/fct, f l = 0.60, respectively. after completion of the cyclic loading a microscopic investigation of the crack characteristics on thin polished sections (5 × 5cm2) has been conducted. table 1 lists the corresponding crack characteristics. compared to the undamaged reference specimens (rel. edyn = 100%), with decreasing relative dynamic e-modulus, the number of cracks increased significantly. at a relative dynamic e-modulus of about 70.2 % the number of cracks increased by more than factor 30. consequently, the total cross section of cracks within the thin polished sections increased similarly. nevertheless, the average crack width and length stayed rather constant at about 5 µm and 1.5 mm, respectively. it can thus be concluded that the observed increased degradation rather resulted from the formation of new cracks than from the expansion of already existing cracks. 3.2. external alkali supply the described microstructural degradation also influenced the penetration behavior of fluids into the concrete. an investigation of the capillary water uptake in concrete was carried out on specimens previously subjected to cyclic loads with varying degrees of damage. therefore, each three karsten-tubes were applied on the surface exposed to tensile stresses. the water uptake was assessed over time. it showed, that the water uptake (more or less merely through capillary suction) occurred faster and higher with increasing damage (figure 4). in a pre-damaged concrete with a relative dynamic e-modulus of approximately 92.7 % the water uptake already increased by about 60 %. at a damage of rel. edynă = ă70.2ă% the water uptake increased by about 250 %. alkali solutions not only penetrate in in-situ concrete pavements through capillary suction, their ingression is in fact considerably increased by the load of the surface traffic. in a test setup that was developed at the ruhr university bochum to specifically explore these effects, the aforementioned concrete beams were arranged in a circular array, to ensure that they could be loaded with six tires consecutively. the top surface of the beams had previously been exposed to flexural tension in the cyclic fourpoint-flexural setup. while the beams were loaded, their top surface was treated with a sodium-chloride solution of 5%-nacl. the tires that were used to apply the load were additionally equipped with a dead weight of up to one ton in order to achieve near to real-life loading conditions for each beam. since the solution was only pressed into the material within the track width of the tires, a merely capillary solution uptake was to be expected for the adjacent areas. this setup thus allowed the comparative assessment of the two types of liquid penetration into the con48 vol. 33/2022 concrete pavements sustainability figure 4. water uptake in concrete with different relative dynamic e-modulus. figure 5. average chloride penetration depth dependent on the number of tire passes and microstructural damage 49 rolf breitenbücher, robin przondziono acta polytechnica ctu proceedings crete [15]. after the overrolling of the large-format concrete beams up to two million times, two small specimens (10 × 10 × 200cm3) were cut from each of the different areas of liquid penetration. the specimens were then split perpendicular to the overrolled surface. the freshly broken surfaces were sprayed with silver nitrate to visualize the chloride penetration depth. the average penetration depths over the cross section of the specimens are visualized in figure 5. the average penetration depth on the specimens with the most damages (rel. edyn = 80.2%) without tire passes amounted to about 23.8 mm already 3ămm above the penetration depth on the undamaged specimens (rel. edyn = 100%). this effect was also observable for the less damaged specimens (rel. edyn = 87.6% and rel. edyn = 93.3%) but less distinctive. the average penetration depth increased only by 0.3 and 0.4ămm respectively. it can further be noticed, that with increasing tire passes, the penetration increases in damaged as well as undamaged specimens. the average penetration depth of an undamaged specimen after 2 million tire passes already amounted to 27.8 mm, whereas for a strongly damaged specimen (rel. edyn = 80.2%) the average penetration depth amounted to 35.2 mm. the influence of the overrolling traffic could not be observed for comparatively little damage (rel. edyn = 93.3%). concludingly, it can be stated, that the influence of the locally induced external pressure through the tire passes is increasing fluid penetration. however, this effect is limited to a depth of about 2 to 4 cm and becomes more significant with increasing microstructural damage. 4. final discussion and conclusion the presented results showed a distinct influence of cyclic mechanical loadings on concrete. due to microstructural degradation, the penetration of liquids into porous concrete is altered and therefore the potential of an intensified damaging asr increases considerably. the dependence of degradation on the number of load cycles and especially on the ratio between the maximum stress acting on the concrete and its bending flexural strength also becomes obvious through the performed ultrasonic measurements. furthermore, the occurred degradation was additionally described through a microscopic analysis of the crack characteristics. previous investigations of the penetration behavior of water towards concrete showed a significant influence of the externally applied pressure on the overall water uptake as well as on the penetration depth. in addition, concrete that had previously been exposed to tensile stresses showed a higher penetration depth, depending on the extent of damage that had been induced before. last but not least, the numerical models that were developed in the context of this research project highly agree with the results of the experimental studies [16, 17]. they can hence be used as a basis for a model, which simulates the entire interaction of influences that may lead to an intensified damaging asr in concrete pavements. acknowledgements the authors would like to thank the german research foundation (dfg) for their financial support of this research project. further, the authors would like to acknowledge the fruitful cooperation with the rub institute for structural mechanics, the federal institute for materials research and testing, berlin, the f.a. fingerinstitute for building materials, weimar and the institute for concrete structures and building materials, karlsruhe. since the foundation of this research group in 2012 outstanding results in the field of asr have been accomplished on the basis of joint projects. references [1] r. springenschmid. die ermittlung der spannungen infolge von schwinden und hydratationswärme im beton. betonund stahlbetonbau 79(10):263-9, 1984. https://doi.org/10.1002/best.198400390. [2] r. breitenbücher, h. ibuk. experimentally based investigations on the degradation-process of concrete under cyclic load. materials and structures 39(7):717-24, 2006. https://doi.org/10.1617/s11527-006-9097-9. [3] h. ibuk. ermüdung von beton unter druckschwellbelastung phd thesis (bochum: ruhr-university bochum), 2008. [4] r. breitenbücher. potenzielle ursachen der rissbildung in betonfahrbahndecken 16. int. baustofftagung ibausil proc. vol 1 (weimar: bauhaus-university weimar), 2006. [5] f. schubert. ausbreitungsverhalten von ultraschallimpulsen in beton und schlussfolgerungen für die zerstörungsfreie prüfung phd thesis (dresden: technical university dresden), 1999. [6] j. o. holmen. fatigue of concrete by constant and variable amplitude loading phd thesis (trondheim: the norwegian institute of technology), 1979. [7] r. springenschmid, w. fleischer. oberflächenrisse in älteren betonfahrbahndecken tiefbau-ingenieurbau-straßenbau, 10:724-32, 1993. [8] f. h. wittmann, p. zhang, t. zhao. eine zerstörungsfreie messmethode zum studium des eindringens von flüssigkeiten in beton 17. int. baustofftagung ibausil proc. 1:969-74, 2009. [9] a. volkwein. untersuchungen über das eindringen von wasser und chlorid in beton phd thesis (munich: tu munich), 1991. [10] m. a. bérubé, j. f. dorion, j. duchesne, et al. laboratory and field investigations of the influence of sodium chloride on alkali-silica reactivity. cement and concrete research 33(1):77-84, 2003. https://doi.org/10.1016/s0008-8846(02)00926-2. [11] j. stark, b. wicht. dauerhaftigkeit von beton, springer verlag, berlin, 2013. 50 https://doi.org/10.1002/best.198400390 https://doi.org/10.1617/s11527-006-9097-9 https://doi.org/10.1016/s0008-8846(02)00926-2 vol. 33/2022 concrete pavements sustainability [12] a. dressler, k. schmidt, d. heinz. avoiding a damaging alkali silica reaction with fly ash, proc. of the int. conf. eurocoalash (copenhagen), 179-85, 2010. [13] c. sievering. dauerhaftigkeit von betonfahrbahndecken unter besonderer berücksichtigung des externen alkalieintrags phd thesis (bochum: institute for building materials, ruhr university bochum), 2012. [14] tl beton-stb 07: technische lieferbedingungen für baustoffe und baustoffgemische für tragschichten mit hydraulischen bindemitteln und fahrbahndecken aus beton korrekturen (forschungsgesellschaft für straßenund verkehrswesen), 2012. [15] r. przondziono, r. breitenbücher. akr unter kombinierten einwirkungen wie beeinflussen zyklische beanspruchungen die degradation und den alkalieintrag in beton tagungsbericht 20. int. baustofftagung ibausil 2:85-92, 2018. [16] r. przondziono, j. j. timothy, m. nguyen, et al. vorschädigungen in beton infolge zyklischer beanspruchungen und deren auswirkung auf transportprozesse im hinblick auf eine schädigende akr. betonund stahlbetonbau 110(1):3-12, 2015. https://doi.org/10.1002/best.201400095. [17] r. przondziono, j. j. timothy, f. weise, et al. degradation in concrete structures due to cyclic loading and its effect on transport processes-experiments and modeling. structural concrete 18(4):519-27, 2017. https://doi.org/10.1002/suco.201600180. 51 https://doi.org/10.1002/best.201400095 https://doi.org/10.1002/suco.201600180 https://doi.org/10.14311/app.2022.33.0624 acta polytechnica ctu proceedings 33:624–630, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague reduction of dissolved oxygen in minimizing corrosion mary grace ventanillaa, jason ongpenga, ∗, nishida takahirob, keiyu kawaaic a de la salle university, department of civil engineering, 2401 taft avenue, 0922 manila, philippines b port and airport research institute, nagase 3-1-1 yokosuka-shi, 239-0826 kanagawa, japan c ehime university, bunkyo-cho 3 matsuyama-shi emime 790-0826, japan ∗ corresponding author: jason.ongpeng@dlsu.edu.ph abstract. concrete structures are susceptible to corrosion especially when exposed to marine environment. in order to minimize the corrosion in reinforced concrete, reduction of dissolved oxygen (do) in mixing water is recommended. the do is the measurement of the amount of free oxygen that is dissolved in water which is proportional to the corrosion rate of steel bars inside concrete. in this paper, the use of industrial waste/by-products, agricultural waste, natural minerals, and green inhibitors as additive for cement in reducing the do level of the mixing water was explored. candidate materials from different types of agricultural waste, industrial waste, natural minerals, and green inhibitors. the percentage difference of do were computed for all types of materials as ash or powder or extract in room temperature. this is the percentage difference of do level in distilled water and the solution mixed with the candidate material having a mass ratio of 3 : 2 and/or 3 : 1. results showed that more than 90% reduction of do level were achieved when mixed with ginger extract, ginger powder, aloe vera extract, ginger pulp, and rice hull ash. future experimental studies using the candidate materials producing reinforced concrete specimens with high reduction of do level in mixing water is recommended. keywords: corrosion, dissolved oxygen, waste. 1. introduction concrete is considered as most extensively used construction material globally and each component has contributed to an environmental impact [1]. it is made up of aggregates (coarse and fine) and paste (cement and water). concrete industry is one of the largest consumers of natural resources, therefore attaining reliable, vigorous, sustainable and economical concrete products has been proposed. cost of concrete products has increased over the years resulting to increase in the shortage of aggregates and worsens situation. use of industrial and agricultural byproducts have gained interests as a potential replacement to concrete manufacturing process [2] since they are rich in silica as well as alumina which are suited as supplementary cementitious material and superior reactivity [3]. through the years, there is scarcity of natural resources. to address such issues, substitution to concretes components could be an answer to environmental impacts, however, substitute materials should impart comparable or enhanced attributes compared to the natural concrete [2]. several factors affect the durability of reinforced concrete structures such as change in temperature that could lead to cracks due to shrinkage, amount of moisture that could permeate concrete, physical factors (strength, resistance to wear and tear, and abrasion), chemical factors like aggressive agents that could interfere with the cement paste and biological factors (organisms) that could damage concrete. one way of achieving durable concrete is to study corrosion which can be defined as "destruction of metal by either chemical, electrochemical or electrolytic reaction within its environment" according to american concrete institute. it is a global problem that deteriorates the durability of concrete structures [5]. also, corrosion of steel in concrete is one of the main reasons for premature failure where in marine environment, chloride ingress is the probable cause for breaking the passivating layer and the onset of extreme corrosion [6]. most marine structures are susceptible to corrosion as it is located near marine environment and may trigger chloride intrusion [7]. nowadays, it is necessary to inspect and alleviate corrosion to extend serviceability of rc structures and is accepted worldwide [4]. three parameters that greatly affect corrosion of reinforcing bars in concrete are shown in figure 1. these are oxygen, ph level, and the permeability of concrete. in this research, reduction of oxygen through dissolved oxygen (do) level in mixing water is the main focus to prevent corrosion inside concrete. additives such as industrial waste/by-products, agricultural waste, natural minerals, and green in624 https://doi.org/10.14311/app.2022.33.0624 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 reduction of dissolved oxygen in minimizing corrosion figure 1. graph showing relationship of embodied carbon for cements by clinker content, source [4] hibitors are considered as candidate materials in reducing the do level of mixing water that could affect concretes resistivity to corrosion. the selection of these materials is based on the result of related studies utilizing them as replacement to cement or aggregate that enhances concrete’s durability. industrial wastes are known to have high ph level which is being measured under varying temperatures to enhance its reactivity [8]. since these wastes no longer have industrial application, converting them to construction materials could improve its properties. in addition, agricultural waste is also found to contain minerals that could improve various properties of concrete. natural minerals as admixture could enhance resistance of concrete to chloride penetration [9]. plant extracts are also considered as these were proven to be good corrosion inhibitors [10]. shown in table 1 is the process of accumulation of the candidate materials from previous researchers. numerous literatures are available considering corrosion in reinforced concrete structures. however, researches associated with oxygen reducing materials haven’t been thoroughly studied. using keyword searches "oxygen", "permeability", "concrete" and "corrosion" resulted to 34 published journals. it is evident that researches concerning on the effects of do in concrete corrosion has an average of only two publications per year and in which approximately one publication per year under cement and concrete research. this study introduces the use of waste materials, plants and natural minerals to contribute significantly in the performance of reinforced concrete, particularly, in the reduction of oxygen that can inhibit corrosion on the steel bars. 2. apparatus and materials 2.1. dissolved oxygen meter sufficient amount of free oxygen at the initial curing stage is critical in the formation of the passive layer protecting the reinforcing steel bars [26]. there is a relevant connection when aggregates in concrete have open porosity as it could increase the permeability of oxygen [27]. corrosion rate of steel is higher on areas with high oxygen contents. measuring the oxygen level of water at a certain location can be used to foresee the performance of different alloys in the same environment [28]. oxygen permeability in reinforced concrete is high using deformed bars which makes it more susceptible to corrosion than plain bars since cracks begin at the ribs of the rebars and eventually becomes surface cracks [29]. do is the measure of the amount of free oxygen that is dissolved in water. factors such as atmospheric pressure, temperature and salinity determine the amount of oxygen that can be dissolved physically in water. rate of corrosion process on the reinforcing steel bars in concrete can be controlled by do on pore solution, thus, reduced do could enhance corrosion resistance [30]. 2.2. candidate materials materials were chosen based on its effect on the performance of concrete. related studies as presented in table 1, utilized the use of these candidate materials as an additive or replacement to either cement and aggregates. table 2 to 4 summarize the candidate 625 m. g. ventanilla, j. ongpeng, n. takahiro, k. kawaai acta polytechnica ctu proceedings corrosion inhibitors candidate materials process of accumulation source wastes industrial waste blast furnace slag acquired from industry as it is a by-product of steel manufacturing [11] bottom ash acquired from industry as it is aby-product of coal power plants [12] fly ash acquired from industry as it is aby-product of coal power plants [13] granite / marble dust acquired from industry producing marble / granite as tiles [14] waste glass powder crushing, sieving using ball mill [15] wood waste ash incineration, combustion [16, 17] agricultural waste coconut husk ash burning under controlled temperature orincineration, pyrolysis [18] rice husk ash burning under controlled temperature or incineration, through annealing furnace [19] corn cob ash burning under controlled temperature [20] sugarcane bagasse ash burning under controlled temperature or incineration, combustion, acquired from sugar mill [13, 17] seashell cleaning, drying, crushing /pulverizing, furnace [21] natural volcanic ash readily available on areas near volcanoes [22] minerals natural zeolite mining [23] plant extracts water hyacinth extraction, drying, grinding [24] aloe extraction [10] ginger extraction, drying [25] table 1. process of accumulation adapted from previous researches. agricultural wastes specimen ash powder3 : 2 3 : 1 3 : 2 3 : 1 bagasse ash bagasse ash aw-ba-2 aw-ba-1 coconut husk ash coconut husk ash aw-cha-2 aw-cha-1 corn cob ash corn cob ash aw-cca-2 aw-cca-1 rice hull ash rice hull ash aw-rha-2 aw-rha-1 shells oyster shell aw-ss-a-os-2 aw-ss-a-os-1 aw-ss-p-os-2 aw-ss-p-os-1 yellow cockles aw-ss-a-yc-2 aw-ss-a-yc-1 aw-ss-p-yc-2 aw-ss-p-yc-1 strombus sicad aw-ss-a-sc-2 aw-ss-a-sc-1 aw-ss-p-sc-2 aw-ss-p-sc-1 brown scallop aw-ss-a-bs-2 aw-ss-a-bs-1 aw-ss-p-bs-2 aw-ss-p-bs-1 clamrose big aw-ss-a-cb-2 aw-ss-a-cb-1 aw-ss-p-cb-2 aw-ss-p-cb-1 granular ark aw-ss-a-ga-2 aw-ss-a-ga-1 aw-ss-p-ga-2 aw-ss-p-ga-1 distant scallop aw-ss-a-ds-2 aw-ss-a-ds-1 aw-ss-p-ds-2 aw-ss-p-ds-1 green mussel aw-ss-a-gm-2 aw-ss-a-gm-1 aw-ss-p-gm-2 aw-ss-p-gm-1 table 2. agricultural wastes as candidate materials in reducing do level. 626 vol. 33/2022 reduction of dissolved oxygen in minimizing corrosion industrial wastes specimen ash powder3 : 2 3 : 1 3 : 2 3 : 1 blast furnace slag blast furnace slag iw-bfs-2 iw-bfas-1 bottom ash bottom ash iw-ba-2 iw-ba-1 fly ash fly ash iw-fa-2 iw-fa-1 marble / granite waste dust marble / granite waste dust iw-wwa-2 iw-wwa-1 table 3. industrial waste as candidate materials in reducing do level. natural mineral specimen ash powder3 : 2 3 : 1 3 : 2 3 : 1 volcanic pumice volcanic pumice nm-vp-2 nm-vp-1 natural zeolite natural zeolite nm-nz-2 nm-nz-1 natural mineral plants specimen extract powder3 : 2 3 : 1 3 : 2 3 : 1 aloe vera aloe vera p-av-2 p-av-1 ginger ginger extract 1 p-ge1-2 p-ge1-1 ginger extract 2 p-ge2-2 p-ge2-1 ginger pulp p-gp-2 p-gp-1 ginger powder p-gpw-2 p-gpw-1 water hyacinth water hyacinth p-wh-2 p-wh-1 table 4. natural and plants as candidate materials in reducing do level. materials based on their ratio with water and type whether powder, ash or extract. a reducibility effect in the do was found using 3 : 1 and 3 : 2 mass ratio (water : specimen ratio), therefore, it was adapted and was considered that it may have a possibility in obtaining the same effect when combined with concrete [31]. candidate materials that were burned, powdered and extracted have composition similar to cement and aggregates that may replace them or be added to enhance concrete properties. usually, those materials that were burned and became ash are associated with cement properties, powdered materials are associated with aggregates while extracted materials may serve as coating to metals to reduce corrosion. in this study, candidate materials are used as an additive to mixing water in reducing do level. figure 2 shows the do meter used and the sample candidate materials. 3. methodology candidate materials were gathered from different sources such as industrial plants, farms, quarries and plantation. industrial wastes were readily available since these are by-products being disposed, whereas, agricultural wastes were burned to form ash which can produce properties similar to cement. moreover, minerals are also readily available as they are part of natural resources while plants, on the other side, were gathered from plantation and were extracted. 3.1. mixing procedure 300ml of distilled water was used to determine the baseline value of do to identify the percentage dif627 m. g. ventanilla, j. ongpeng, n. takahiro, k. kawaai acta polytechnica ctu proceedings figure 2. do meter and sample candidate materials. figure 3. candidate materials for do testing. ference of do when mixed with the candidate materials. do of water was found out to be 5.08ppm. do of the mixture were measured by placing 100grams and 200grams of materials into a bottle with 300ml of distilled water (3 : 1 and 3 : 2 ratio), stirred until thoroughly mixed to form a solution. at room temperature (average 26 degrees celsius), do level was measured using do meter. except for the case of garlic powder and rha, the ratio used was 10 : 1 since these materials were observed to be low density, thus, bagasse ash and wood waste ash were tested only in 3 : 1 ratio. figure 3 shows how the specimens were prepared prior to do testing. 3.2. percentage difference of do the equation used to determine the do is shown below. %dodif f = dowater − dosolution dowater · 100 (1) where: • dowater = do level of distilled water at room temperature • dosolution = do level of mixed solution at room temperature 3.3. results and discussion shown in table 5 are the materials with highest do level reduction represented by the percentage difference resulted to a reduction of more than 90% which include ginger extract, ginger powder, aloe vera, ginger pulp, and rice hull ash. materials do diff.(ppm or mg/l) % do diff p-ge-1 4.99 98.23 p-ge-2 4.98 98.03 p-gpw 4.97 97.83 p-av-1 4.96 97.64 p-av-2 4.91 96.65 p-gp-1 4.81 94.69 aw-rha 4.77 93.90 table 5. top percentage do difference of the candidate materials. 628 vol. 33/2022 reduction of dissolved oxygen in minimizing corrosion 4. conclusion corrosion inside concrete is very complex with a lot of factors contributing to it. in this study, the objective is to reduce the do level of mixing water that can reduce the corrosion of steel rebars inside concrete. identification of candidate materials from agricultural waste, industrial waste, minerals, and plants were made in the region. among the candidate materials, it can be observed that plants such as ginger and aloe extracts have the highest capacity to reduce oxygen in a solution due to its inherent properties and constituents that may inhibit corrosion followed by agricultural and industrial wastes. future studies are recommended to test the candidate materials as an additive to mixing water in concrete by conducting permeability test and evaluating through scanning electron microscope sem, to identify its efficiency as part of the binding material in the reduction of corrosion. references [1] b. bhardwaj, p. kumar. waste foundry sand in concrete: a review. construction and building materials 156:661-74, 2017. https: //doi.org/10.1016/j.conbuildmat.2017.09.010. 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[31] t. nishida, n. otsuki, a. saito, et al., editors. influence of reducibility environment around steel bars on steel corrosion in concrete. proceedings of the seminar workshop on the utilization of waste materials e, 2017. 630 https://doi.org/10.1016/j.conbuildmat.2013.02.069 https://doi.org/10.1016/j.jobe.2018.01.002 https://doi.org/10.1016/j.jclepro.2018.12.299 https://doi.org/10.1016/j.cemconcomp.2015.10.022 https://doi.org/10.1016/j.conbuildmat.2017.01.075 https://doi.org/10.1016/s0008-8846(01)00482-3 https://doi.org/10.1016/j.cemconres.2019.105822 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0076 acta polytechnica ctu proceedings 40:76–82, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague extracting general knowledge of model parameters for clays out of numerous laboratory tests veronika pavelcováa, ∗, tomáš jandab a czech technical university in prague, faculty of civil engineering, department of geotechnics, thákurova 7, 166 29 prague, czech republic b czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague, czech republic ∗ corresponding author: veronika.pavelcova@fsv.cvut.cz abstract. this paper is concerned with the process of determining the material parameters of advanced constitutive models for clays. provided results are applicable already in the early stages of geotechnical survey when no detailed data from laboratory tests of soils expected at construction site are known. a worldwide database of laboratory tests obtained during four year operation of web calibration application excalibre was used for this purpose. this application allows automatic calibration of three advanced constitutive soil models based on detailed data from triaxial and edometric tests. the paper describes the process of determining the confidence intervals of each material parameter of modified cam-clay and hypoplastic clay models for low (cl) and high (ch) plasticity clay according to unified soil classification system (uscs) and tabulates their typical ranges. the purpose of presented work was to contribute to the use of advanced constitutive models among practical engineers. keywords: hypoplastic clay, modified cam-clay, calibration, material parameters. 1. introduction in engineering practice, it can be observed that majority of geotechnical engineers limit themselves to the use of basic non-linear soil models of the mohrcoulomb or drucker-prager type when describing the interaction of the structure and the subsoil and mostly avoid critical state models when solving standard tasks. the general lack of knowledge of these advanced material models among practical engineers may be one of the reasons. also standard civil engineering studies and the most commercial geotechnical programs primarily promote only basic models. the second reason is that it is significantly more difficult to determine the input parameters of advanced non-linear models in comparison to the basic linear models. while the inputs for the use of basic nonlinear models are shear strength parameters that can be determined directly from the results of standard laboratory or field tests, and their ranges are even tabulated for each soil type, critical state models require prescription of parameters describing the behavior of the given material not only under isotropic compression but also parameters characterizing the current state of the soil. a practical engineer often encounters a situation where, although he knows that there exist models which are capable to describe the behavior of modeled soil significantly more accurately than linear constitutive models, he is no longer able to determine the necessary parameters that govern the model. 2. critical state models both, the hypoplastic clay [1] and modified camclay [1, 2], constitutive models are based on the theory of critical state. the main features of these models are similar but unlike the hypoplastic model the modified cam-clay model distinguishes the elastic and plastic part of strain components and therefore allows the straightforward visualization of permanently strained plastic zones in fem calculations. on the other hand, the model of hypoplastic clay can correctly represent the associated gradual reduction in stiffness upon activation of plastic behavior rather than a sudden drop as predicted by the modified cam-clay model. 2.1. hypoplastic clay the hypoplastic model for clay takes into account the following important properties of soils: • the stiffness depends on the current value of the mean effective stress σeffm and the direction of loading. • the strength depends on the value of the mean effective stress σeffm and current density characterized by the void ratio e. the model distinguishes between material parameters that are constant for all possible states of a particular soil and state variables that evolve during straining. in this sense the hypoplastic model is fundamentally different from the mohr-coulomb model which employs distinct values of its material parameters for the same soil in different density states. in 76 https://doi.org/10.14311/app.2023.40.0076 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 40/2023 extracting knowledge of model parameters out of laboratory tests parameter name unit material parameters n position of the normal consolidation line [-] λ∗ slope of the normal consolidation line [-] κ∗ slope of the swelling line [-] φc friction angle in the critical state [°] ν parameter controlling the value of the shear modulus [-] state parameter e void ratio [-] table 1. material parameters of hypoplastic clay model. figure 1. normal consolidation line and swelling line assumed in hypoplastic model for clay [1]. particular, the hypoplastic model for clay requires five material parameters and one state variable summarized in table 1. 2.1.1. parameters n and λ∗ parameters n and λ∗ define the position and the slope of the isotropic normal compression line (ncl) in the ln(σeffm )×ln(1+e) space as illustrated in figure 1. the normal compression line defines a theoretical response of the soil sample during isotropic compression test. since the model assumes incompressible grains the resulting volumetric strain depends on the decrease of volume of pores only and thus on the void ratio. 2.1.2. parameter κ∗ parameter κ∗ defines the slope of the swelling line in the ln(σeffm ) × ln(1 + e) space as illustrated in figure 1. unlike the ncl which is fixed by parameter n the position of the swelling line depends on the previous compression stress. in addition, unlike simple elastoplastic models, the slope of the unloading-reloading line is non-linear in hypoplasticity, so that the paparameter name unit material parameters e0 position of the normal consolidation line [-] λ slope of the normal consolidation line [-] κ slope of the swelling line [-] mcs slope of the critical state line [-] ν poisson’s ratio [-] state parameter pc the preconsolidation pressure [pa] table 2. material parameters of modified cam-clay model. rameter κ∗ should thus be calibrated by means of simulation of a laboratory test, rather than by directly evaluating its slope. 2.1.3. parameter φc parameter φc is the critical angle of internal friction. the angle of repose can not be measured for finegrained soils and the critical friction angle has to be determined from triaxial shear test. the critical friction angle correspond to the maximal mobilized friction angle for normally consolidated samples where no softening occurs during the triaxial shear test. or the critical friction angle corresponds to the mobilized friction angle after the peak for overconsolidated samples. 2.1.4. parameter ν parameter ν controls the value of the shear modulus. the bulk and shear modulus are linked through the parameter ν in the isotropic normally consolidated state. during shear test the shear modulus evolves non-linearly and its magnitude is being controlled by the parameter ν and by the distance from the failure state. 2.2. modified cam-clay the modified cam-clay takes into account the following important properties of soils: • the stiffness depends on the current value of the mean effective stress σeffm and the direction of loading. • the strength depends on the value of the mean effective stress σeffm a current density characterized by the void ratio e or by the preconsolidation pressure pc. the model introduces five material parameters that are constant for all possible states of a particular soil and one state variable that evolves during straining. all parameters are summarized in table 2. 77 veronika pavelcová, tomáš janda acta polytechnica ctu proceedings figure 2. normal consolidation line and swelling line assumed in modified cam-clay model [2]. 2.2.1. parameters e0 and λ parameters e0 and λ define the position and the slope of the isotropic normal compression line (ncl) in the ln(σeffm ) × e space as illustrated in figure 2. the normal compression line defines a theoretical response of the soil sample during isotropic compression test. since the model assumes incompressible grains the resulting volumetric strain depends on the decrease of volume of pores only and thus on the void ratio. 2.2.2. parameter κ parameter κ defines the slope of the unloading/reloading line in the ln(σeffm )×e space as illustrated in figure 2. unlike the ncl which is fixed vertically by parameter e0 the position of the swelling (unloading/reloading) line depends on the maximal previous compression stress denoted as the preconsolidation pressure pc. 2.2.3. parameter mcs parameter mcs is the slope of the critical state line plotted in the σeffm × j space, see figure 3. the value is computed from the critical state angle φc by: mcs = 2 √ 3 sin φc 3 − sin φc . (1) 2.2.4. parameter ν parameter ν is the poisson ratio known from linear elasticity. it controls the ratio of radial and axial elastic strains under uniaxial loading and consequently the ratio between the stress dependent bulk and shear modulus. 2.2.5. preconsolidation pressure pc the model has one state variable – the preconsolidation pressure pc. for an isotropic compression test the preconsolidation pressure pc corresponds to the maximal mean stress σeffm to which the soil has been subjected. the kinematics of the model is evident from figure 2. soil sample with an initial state a (it has void ratio ea and preconsolidation pressure pac ) is loaded isotropically up to the compressive mean stress σeffm = pac . the deformation during this loading is nonlinear but elastic, i.e. if the soil is unloaded from this point it returns along the κ-line back to the state with e = ea and no permanent volumetric strain occurs. if the soil is loaded beyond the point pac it moves along the λ-line. during this loading both the elastic and the plastic volumetric strain develop. when the soil is unloaded from the point σeffm = pbc it moves along the new κ-line defined by the new value of preconsolidation pressure σeffm = pbc. as evident from the initial and final (in unloaded state) values of the void ratio ea > eb the permanent volumetric strain developed during this loading sequence [1]. 3. automated calibration of constitutive models even though the development of more advanced constitutive soil model theories which take into account bounding surface plasticity, whether those based on the theory of plasticity or hypoplasticity [3, 4], or barodesy [5] continues steadily forward, the development of automated calibration procedures stays behind. the calibration procedures are regarded as the inverse analysis, since the parameters of a constitutive model are not known in advance but the reaction of the system is known. the goal of the calibration is to minimize the objective error function e(u ) which is defined as a function of a difference between the observation or experiment uexp and simulation unum. the most common is the least square method. nevertheless, more sophisticated dimensionless formulas can be used [6]. 3.1. excalibre as mentioned in the introduction, the faculty of civil engineering of ctu in cooperation with the faculty of science of charles university developed a excalibre web application [7] which performs automatic deterministic calibration of the hypoplastic model for sands [8], hypoplastic model for clays [9] and modified cam-clay model [10]. from 2018, the application [1] is available free of charge for registered users who agree to provide the uploaded data for research and development purposes. a typical workflow of model calibration in excalibre is the following: (1.) the user downloads a laboratory protocol template which is prepared for the user in the form of an ms excel file. the template is different for clays and sands. (2.) the user fills in the results of laboratory tests. a minimum of one consolidated isotropically undrained (ciup) triaxial test with pore pressure measurement for clay soils or one consolidated isotropically drained (cid) triaxial test for sandy 78 vol. 40/2023 extracting knowledge of model parameters out of laboratory tests figure 3. yield surface of the modified cam-clay model [2]. soils and one standard oedometric (oed) test must be completed for successful calibration. in addition to those already mentioned, it is possible to fill in the classification according to uscs and index characteristics of the soil. these parameters are currently not used in calibration but they are stored for research purposes. (3.) the user uploads the laboratory protocol back to the web application. as a result, the application proposes optimal model parameters and graphs comparing the course of measured and simulated laboratory tests. these graphs serve to visually check the calibration of the model. (4.) it is possible to manually modify individual model parameters to adjust the calibration or to examine the influence of individual parameters on model predictions after the calibration. large pool of available data collected worldwide propose an effective basis for statistical processing, determination of typical ranges of individual parameters and search for correlations between individual parameters presented in this paper. 4. data processing all data work was concentrated in universal scripts written in the python programming language due to the repeatability of laboratory test data analysis. it is a higher-level programming language with dynamic support for data types and support for programming paradigms. the programming language [11] is freely available. the format conversion of laboratory protocols from microsoft excel spreadsheet xlsx excalibre template to comma-separated values csv was the first step of data processing. it was a necessary step due to the complexity of the xlsx file structure. xlsx file is a zip package that contains xml files representing individual sheets of the spreadsheet which are always supported by rels files containing specifications for linking individual xml files together to form a complete file. compared to complex xlsx file, csv is a simple file format that is intended for exchanging tabular data between different systems. the data of the csv file are stored in rows just like in the default table while the individual items in the row are further separated by a predetermined character. the individual sheets of the input xlsx laboratory protocol form separate csv files which are, for organizational reasons, stored in a folder with a name corresponding to the original name of the xlsx file so that one folder represents one complete laboratory protocol. written script cycles through the folders representing individual laboratory protocols with csv files replacing microsoft excel spreadsheet and extracts characteristic information about the laboratory protocol using predefined classes. the extracted data is further inserted into a dictionary-type data structure. a dictionary is a data type sometimes called an associative array. it consists of key-value pairs. each key-value pair maps a key to an associated value. in the algorithm described , the key corresponds to the name of a folder representing one laboratory protocol. key values are assigned using classes so that one line of the dictionary represents the data of one laboratory protocol. the first value of the dictionary corresponds to the relative path to the file, assigned attributes are: uscs classification of soil, basic index characteristics such as specific gravity, atterberg’s limits or angle of repose. it is followed by representation of particle size distribution and loading and quantity of performed oedometric and triaxial tests. the excalibre application contains a public laboratory test database so that these protocols can be used for testing the functionality of the calibration application by a new user. knowing that, filtering of duplicate laboratory protocols was another step of data processing. a simple condition consisting of • uscs classification of soil, • the largest sieve and the corresponding passing in the grain size analysis, • the initial void ratio of the oedometric test of the naturally consolidated soil sample, • the initial void ratio of the oedometric test of the reconstituted soil sample 79 veronika pavelcová, tomáš janda acta polytechnica ctu proceedings hypoplastic clay modified cam-clay λ∗ κ∗ n ν φc λ κ e0 ν mcs count 7 7 7 7 7 7 7 7 7 7 mean value 0.119 0.010 1.396 0.276 25.3 0.297 0.016 2.784 0.307 0.997 standard deviation 0.082 0.003 0.659 0.062 2.0 0.365 0.012 2.521 0.065 0.088 minimum value 0.063 0.006 0.872 0.170 22.8 0.104 0.010 1.237 0.220 0.888 25 % quantile 0.068 0.009 0.960 0.245 23.9 0.107 0.010 1.383 0.250 0.935 50 % quantile 0.094 0.010 1.173 0.280 25.6 0.155 0.010 1.752 0.340 1.008 75 % quantile 0.123 0.010 1.521 0.320 26.2 0.249 0.018 2.734 0.355 1.032 maximum value 0.296 0.014 2.763 0.350 28.8 1.107 0.039 8.263 0.380 1.147 table 3. basic statistical description of calibrated model parameters for ch soil. was used for the determination and removal of duplicate laboratory protocols. the results of this classdriven filtering revealed that only 58 of the total 1 924 laboratory protocols are non-duplicated. the remaining protocols are duplicates of another 161 protocols in various frequencies. in total, the database would therefore contains 219 unique laboratory protocols. considering uscs composition of non-duplicate protocols only low (cl) and high (ch) plasticity clay soils were selected for further processing. all of the laboratory protocols of soils classified as ch or cl have been uploaded to the excalibre web application and material parameters for hypoplastic model for clay soils and modified cam-clay were calibrated and added to the dictionary-like dataframe collecting protocol information. thus the already existing dataframe was extended by 5 columns with material parameters of hypoplastic clay model (see table 1) and another 5 columns for modified cam-clay model material parameters (see table 2). 5. results the database contained only 7 laboratory protocols of soil classified as high-plasticity clay after removing all duplicates. all of them were successfully calibrated for both of constitutive models – hypoplastic clay and modified cam-clay. the summarization of basic statistical description of calibrated model parameters is summed up in table 3. strong correlations between atterberg’s limits (liquid limit plastic limit wl and plastic limit wp ) and slope of the ncl line of both calibrated models have been observed. example of correlation between plastic limit and slope of the ncl assumed in hypoplastic model for clay can be seen in figure 4 and the same trend applies to the position of the ncl. all considerable correlations are summarized in table 4 but limited number of protocols has to be considered. overall 72 laboratory protocols of soil classified as low-plasticity clay contained the database after removing all duplicates of which it was possible to calibrate 56 for hypoplastic clay and 61 for modified cam-clay model. the reason why not all of the protocols were successfully calibrated will be further examined. basic statistical description of calibrated model parameters 25 30 35 40 plastic limit wp [-] 0.10 0.15 0.20 0.25 0.30 sl op e of th e no rm al c on so lid at io n lin e * [ -] figure 4. correlation between plastic limit and slope of the ncl assumed in hypoplastic model for clay. hypoplastic clay mod. cam-clay λ∗ n ν φc λ e0 mcs wl 0.92 0.95 0.41 0.91 0.94 0.40 wp 0.95 0.92 0.52 0.44 0.89 0.88 0.44 table 4. correlations between atterberg’s limits of high-plasticity clay soil and model’s parameters. is also summarized and can be examined in table 5. examples of parameter’s distribution are shown in histogram in figure 5. in comparison to high-plasticity clay, only moderate correlations between atterberg’s limits and model parameters have been observed for low-plasticity clay. in particular, higher soil moisture at atterberg’s limits indicates higher parameter n with correlation coefficients about 0.41. the similar trend also applies to modified cam-clay’s parameter e0 with correlation coefficients about 0.46 for the liquid limit and the 0.41 for plastic limit. correlation 0.40 has been found between moisture at liquid limit and the slope of the ncl line λ. another correlations seem insignificant for practical applications. 6. conclusions this paper briefly explained each individual input parameters of both hypoplastic and modified cam80 vol. 40/2023 extracting knowledge of model parameters out of laboratory tests hypoplastic clay modified cam-clay λ∗ κ∗ n ν φc λ κ e0 ν mcs count 56 56 56 56 56 61 61 61 61 61 mean value 0.067 0.008 0.948 0.217 29.1 0.107 0.010 1.388 0.270 1.120 standard deviation 0.035 0.005 0.283 0.109 6.6 0.062 0.006 0.597 0.112 0.320 minimum value 0.001 0.000 0.507 0.010 2.9 0.002 0.000 0.112 0.010 0.101 25 % quantile 0.042 0.005 0.725 0.155 25.0 0.059 0.008 0.949 0.200 0.984 50 % quantile 0.064 0.007 0.899 0.230 30.0 0.096 0.010 1.246 0.260 1.199 75 % quantile 0.090 0.010 1.170 0.290 33.3 0.149 0.011 1.797 0.390 1.274 maximum value 0.178 0.025 1.667 0.400 43.6 0.284 0.036 3.075 0.400 1.790 table 5. basic statistical description of calibrated model parameters for cl soil. 0.00 0.05 0.10 0.15 slope of the normal consolidation line * [-] 0 2 4 6 8 10 12 14 16 18 fr eq ue nc y (a). hypoplastic clay model. 0.00 0.05 0.10 0.15 0.20 0.25 slope of the normal consolidation line [-] 0 2 4 6 8 10 12 14 16 fr eq ue nc y (b). modified cam-clay model. figure 5. histograms of slopes of the ncl for hypoplastic clay and modified cam-clay model. clay constitutive model and it summarized the issue of automatic parameter calibration. excalibre web calibration application, a practical tool for calibrating these material parameters based on the results of standard laboratory tests, was introduced in section 2 and section 3. then a method of how could existing data be used to obtain general knowledge of the properties of these individual parameters was presented in section 4 and finally, the paper summarized extracted basic statistical data of model parameters with respect to uscs classification in section 5. the main aim of this research is to help with development of a robust automatic calibration procedure which will be applicable at any stage of geotechnical survey to promote the use of advanced models in geotechnical practice. the determination of the confidence intervals along with correlations between material parameters and index soil properties was only the first step of data analysis and other more complex studies will follow. acknowledgements the support provided by the sgs project no. sgs22/030/ohk1/1t/11 and the gačr project no. 22-12178s is gratefully acknowledged. references [1] excalibre – soilmodels automatic calibration. [2022-07-26]. https://soilmodels.com/excalibre-en/ [2] m. šejnoha, j. pruška, t. janda, m. brouček. metoda konečných prvků v geomechanice: teoretické základy a inženýrské aplikace. české vysoké učení technické, 2015. isbn 978–80–01–05743–8. [3] j. jerman, d. mašín. hypoplastic and viscohypoplastic models for soft clays with strength anisotropy. international journal for numerical and analytical methods in geomechanics 44(10):1396–1416, 2020. https://doi.org/10.1002/nag.3068 [4] d. mašín. clay hypoplasticity model including stiffness anisotropy. géotechnique 64(3):232–238, 2014. https://doi.org/10.1680/geot.13.p.065 [5] d. kolymbas. introduction to barodesy. géotechnique 65(1):52–65, 2015. https://doi.org/10.1680/geot.14.p.151 [6] t. kadlíček. parameters identification of advanced constitutive models of soils. ph.d. thesis, czech technical university in prague, 2019. [7] t. kadlíček, t. janda, m. šejnoha, et al. automated calibration of advanced soil constitutive models. part ii: hypoplastic clay and modified cam-clay. acta geotechnica 17:3439–3462, 2022. https://doi.org/10.1007/s11440-021-01435-y [8] p.-a. von wolffersdorff. a hypoplastic relation for granular materials with a predefined limit state surface. mechanics of cohesive-frictional materials: an international journal on experiments, modelling and computation of materials and structures 1(3):251–271, 1996. https://doi.org/10.1002/(sici)10991484(199607)1:3<251::aid-cfm13>3.0.co;2-3 81 https://soilmodels.com/excalibre-en/ https://doi.org/10.1002/nag.3068 https://doi.org/10.1680/geot.13.p.065 https://doi.org/10.1680/geot.14.p.151 https://doi.org/10.1007/s11440-021-01435-y https://doi.org/10.1002/(sici)1099-1484(199607)1:3<251::aid-cfm13>3.0.co;2-3 https://doi.org/10.1002/(sici)1099-1484(199607)1:3<251::aid-cfm13>3.0.co;2-3 veronika pavelcová, tomáš janda acta polytechnica ctu proceedings [9] d. mašín. clay hypoplasticity with explicitly defined asymptotic states. acta geotechnica 8(5):481–496, 2013. https://doi.org/10.1007/s11440-012-0199-y [10] k. roscoe, j. burland. on the generalized stressstrain behaviour of wet clay. in engineering plasticity, pp. 235–609. cambridge university press, 1968. [11] python. [2022-07-26]. https://www.python.org 82 https://doi.org/10.1007/s11440-012-0199-y https://www.python.org acta polytechnica ctu proceedings 40:76–82, 2023 1 introduction 2 critical state models 2.1 hypoplastic clay 2.1.1 parameters n and lambda* 2.1.2 parameter kappa* 2.1.3 parameter phi c 2.1.4 parameter nu 2.2 modified cam-clay 2.2.1 parameters e0 and lambda 2.2.2 parameter kappa 2.2.3 parameter mcs 2.2.4 parameter nu 2.2.5 preconsolidation pressure pc 3 automated calibration of constitutive models 3.1 excalibre 4 data processing 5 results 6 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0219 acta polytechnica ctu proceedings 33:219–225, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the role of conceptualization in the evaluation of sustainable concrete michael henrya, ∗, joel oponb a shibaura institute of technology, toyosu 3-7-5, koto-ku, tokyo, japan b msu iligan institute of technology, iligan city 9200, philippines ∗ corresponding author: mwhenry@shibaura-it.ac.jp abstract. improvement of the sustainability of concrete materials will be realized through the development of analytical tools that facilitate sustainable design and evaluation. however, these processes may be dependent on how sustainability is conceptualized for concrete. conceptualization is the means by which sustainability is operationalized by creating a structure that connects a qualitative goal to its quantitative indicators. as there exists no established definition of sustainability for the concrete field, conceptualization is a source of uncertainty in the sustainability evaluation of concrete. this paper explores the role conceptualization plays in the evaluation of concrete material sustainability by analyzing its effects using multicriteria analysis and a sustainability indicator framework to quantify sustainability for concrete materials. six analytical scenarios are explored using frameworks based on direct loading, the three pillars of sustainability, and the sustainable development goals, together with two aggregation methods. it was found that the most sustainable concrete mix varied by scenario, but one concrete mix combining blast furnace slag and high grade recycled aggregates could be judged as the most sustainable due to its highest mean score and lowest variance across all analytical scenarios, which suggests it as the mix least sensitive to methodological choices on conceptualization and aggregation. overall, however, the sustainability scores were highly correlated between the different scenarios. keywords: conceptualization, multi-criteria analysis, sustainable design. 1. introduction the use of concrete incurs significant impacts to environmental systems, including climate change, ecosystem degradation, resources depletion, and pollution [1]. to address these issues, it is necessary to adopt more sustainable materials in construction, which requires evaluation of the sustainability of concrete considering the multi-dimensional nature of sustainable development. literature presents a variety of indicators and evaluation methods for tackling the sustainability evaluation of concrete [2]. examples include comparative analysis of the trade-off between environmental impact and other performances, such as safety and cost [3], and the application of life cycle assessment (lca) to quantify the impacts across the life cycle [4]. multicriteria decision-making techniques, such as analytical hierarchy process, have also been explored to identify the most sustainable construction material [5]. however, these approaches are non-equivalent and lead to uncertainty in the evaluation result, as differing conclusions may be reached when adopting different methods. the lack of a formalized conceptual paradigm for concrete sustainability may contribute to the multiplicity of evaluation methods, which introduces uncertainty into the sustainability evaluation. conceptualization is the process whereby a complex system is decomposed into increasingly smaller components until they can be measured by individual indicators. this creates a structure that rationally links quantitative measures to a qualitative objective. one of the most well-known conceptualizations of sustainable development is the "three pillars" model, wherein the nebulous goal of "sustainability" is first decomposed into the three dimensions of society, environment, and economy. these dimensions are then further broken down into relevant themes and sub-themes, which are ultimately measured by specific indicators [6]. the three pillars, however, represent just one conceptual paradigm for sustainability, and there exist other, equally valid conceptualizations derived from alternative perspectives on sustainable development [7]. formalizing a conceptual framework for concrete sustainability is essential to unify sustainability evaluation for the concrete field. in this paper, the effect of conceptualization on the sustainability evaluation of concrete materials is explored through an exploratory evaluation using multicriteria analysis. three conceptual frameworks for sustainable development are examined: direct loading, wherein sustainability is measured directly from individual indicators; the aforementioned three pillars paradigm; and the sustainable development goals (sdgs), which represent the latest global framework for pursuing sustainable development through a set of 17 goals. compensability, or the treatment of trade-offs between indicators, pillars, or sdgs, is also examined. 219 https://doi.org/10.14311/app.2022.33.0219 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en michael henry, joel opon acta polytechnica ctu proceedings � )ljxuh����6hwxs�iru�vxvwdlqdelolw\�hydoxdwlrq�e\�0&$�figure 1. setup for sustainability evaluation by mca. mix opc bfs fa w ng rg s w/b bfs/b fa/b f ′ c−28 [kg/m3] [−] [−] [−] [mpa] opc-ng 325 0 0 184 1063 0 783 0.57 − − 30.6 opc-rg 325 0 0 184 0 1063 783 0.57 − − 30.6 fa-ng 238 0 94 187 950 0 847 0.56 − 0.28 30.2 fa-rg 238 0 94 187 0 950 847 0.56 − 0.28 30.2 bfs-ng 200 133 0 192 965 0 806 0.58 0.40 − 30.4 bfs-rg 200 133 0 192 0 965 806 0.58 0.40 − 30.4 opc: ordinary portland cement, bfs: blast furnace slag, fa: fly ash, w: water, ng: normal coarse aggregate, rg: high grade recycled coarse aggregate, s: fine aggregate, b: total binder materials table 1. mix proportions and 28-day compressive strengths of the concrete mixes. the results and discussion are expected to provide insights into how to address uncertainty in conceptualization for improving the transparency and robustness of sustainability evaluation for concrete. 2. methodology 2.1. multicriteria analysis setup sustainability evaluation was carried out using multicriteria analysis (mca). mca is widely utilized to support decision-making by identifying the most sustainable option among a set of potential alternatives [8]. the mix proportions of a set of concrete mixes were used as the inputs. the conventional stages of mca are: indicator selection (i), data normalization (n), weighting assignment (w), and aggregation (a). the output of this process is typically a sustainability score (s) or rank (r) for each alternative, which can then be used for quantitative comparison and decision-making. in this paper, the selection of conceptual framework (c) for linking the indicators to sustainability is treated as an additional stage in the mca process. the setup for sustainability evaluation by mca is illustrated in figure 1. only a single indicator set, normalization method, and weighting scheme were adopted for their respective stages. however, three conceptual frameworks and two aggregation methods were utilized for their stages for a total of six analytical scenarios leading to six sustainability scores for each mix alternative. these scores are then statistically examined to explore how a plurality of perspectives on concrete sustainability may affect the evaluation result. 2.2. input concrete mix data the concrete mix alternatives were sampled from a database assembled by noguchi et al. [9]. a target compressive strength of 30 mpa at 28 days was set, and three concrete mixes with different binder contents were chosen. an additional three mixes were then generated assuming 100% replacement of normal coarse aggregates with high grade (class h) recycled coarse aggregates. the japan industrial standard (jis a 5021) indicates that high grade recycled aggregates should be equivalent to normal aggregates, so the concrete compressive strengths are assumed to remain unchanged. however, the energy consumption and emissions footprints of high grade recycled aggregates are higher than that of normal aggregates due to the recycling process, so the inclusion of these mixes will explore the tradeoff between increased consumption of recycled materials and increased environmental impacts in the sustainability evaluation. the mix proportions are summarized in table 1. 2.3. sustainability indicators a subset of 14 quantitative indicators was adopted from a comprehensive set of sustainable concrete material indicators (scmis) [2], with descriptions and characteristics given in table 2. measurement was carried out using available inventory data from the japan society of civil engineers (for ng, rg, s) and the life cycle assessment society of japan (for opc, bfs, fa), characterization values from the 220 vol. 33/2022 evaluation of sustainable concrete scmi id description unit loading by concept three pillars sdgs 1 energy consumption, renewable and non-renewable mj all 7 2 raw materials consumption, primary and secondary kg all 12 3 water consumption kg all 6 4 recycled materials consumption, primary and secondary kg all 12 5 co2 emissions kg-co2 env 9 6 sox emissions kg-sox envsoc 11 7 nox emissions kg-nox envsoc 11 8 particulate matter emissions (pm) kg-pm env 11 28 global warming potential (gwp) kg-co2 eq env 13 29 photochemical ozone creation potential (pocp) kg-c2h4 eq env 11 30.2 acidification potential, aquatic (ap) kg-so2 eq env 14 31.1 eutrophication potential, terrestrial (ep) kg-po4 eq env 2 34 human toxicity potential (hp) kg-c6h4cl2 eq soc 3 40 material cost of concrete yen eco 8 all: environmental, social, and economic, envsoc: environmental and social, env: environmental, soc: social, eco: economic table 2. summary of the adopted sustainable concrete material indicators. mix scmi 1 scmi 2 scmi 3 scmi 4 scmi 5 scmi 6 scmi 7energy raw mats. water recyc. mats. co2 sox nox opc-ng 1213.0 2225.8 184.0 123.8 255.1 0.053 0.513 opc-rg 1563.8 1162.8 184.0 1186.8 270.9 0.053 0.539 fa-ng 956.6 2075.1 186.7 184.8 190.2 0.043 0.379 fa-rg 1270.1 1125.2 186.7 1134.7 204.3 0.043 0.402 bfs-ng 861.8 2005.3 192.0 209.2 162.6 0.038 0.320 bfs-rg 1180.4 1039.9 192.0 1174.6 176.9 0.039 0.344 mix scmi 8 scmi 28 scmi 29 scmi 30.2 scmi 31.1 scmi 34 scmi 40pm gwp pocp ap ep htp cost opc-ng 0.015 255.1 0.017 0.412 0.067 0.620 5775 opc-rg 0.016 270.9 0.018 0.430 0.070 0.652 5775 fa-ng 0.012 190.2 0.013 0.308 0.049 0.458 5266 fa-rg 0.012 204.3 0.013 0.324 0.052 0.487 5266 bfs-ng 0.010 162.6 0.011 0.262 0.042 0.388 5679 bfs-rg 0.011 176.9 0.011 0.279 0.045 0.416 5679 table 3. mix proportions and 28-day compressive strengths of the concrete mixes. lca database of leiden university (for gwp, pocp, ap, ep, hp), and constituent material costs reported by henry et al [10]. the cost of high grade recycled aggregates was assumed the same as that of normal aggregates due to lack of data. the raw indicator values for the six mixes were calculated as shown table 3. the loadings of the indicators within the three pillars and sdgs concepts, as established by opon and henry [2], are also shown in table 2. the environmental pillar is most represented among the three pillars, with 12 of the 14 indicators loading to the environment. ten of the 17 goals are evaluated under the sdg concept, with sdg 11 (sustainable cities and communities) receiving the most representation with four indicators highly relevant to this goal. 2.4. normalization method as the adopted indicators exhibit differing characteristics and behavior, their data were normalized by standardization. standardization was carried out through a combination of z-scores, which convert the data based on a mean of zero and standard deviation of one (eq. 1), and t-scores, which shift the data to a non-negative scale of zero to 100 (eq. 2). z = x − µ s (1) t = (z × 10) + 50 (2) where z: z-score, t: t-score, x: raw value, µ: average, and s: standard deviation,. in some cases, the t-score is divided by 100 to scale the data from zero to one. 221 michael henry, joel opon acta polytechnica ctu proceedings �d��'luhfw�ordglqj� �e��7kuhh�slooduv�frqfhswxdo�iudphzrun� �f��6'*v�frqfhswxdo�iudphzrun� figure 2. conceptual frameworks for sustainability with indicator loadings. 2.5. conceptual frameworks to represent different conceptual approaches to sustainability, three frameworks linking the individual sustainability indicators to the overall objective of concrete sustainability were constructed (figure 2). the first framework, direct loading, assumes no priority areas or sub-categories; that is, all indicators load directly to the sustainability performance of the concrete materials. the second framework decomposes concrete sustainability into the three pillars of society, environment, and economy, and each indicator then loads to its related pillar, or pillars (for multidimensional indicators), following the loadings given in table 2. the final framework sets the ten relevant sdgs as the intermediate priority areas, with each indicator loading to its respective sdg. 2.6. weighting scheme as no information was available to establish the comparative importance of the indicators, pillars, and sdgs, equal weighting was applied for all loadings between levels of the conceptual framework. this is consistent with the united nations resolution adopting the sdgs, which states that they will give "equal priority" to all goals in their implementation efforts [11]. 2.7. aggregation methods two aggregation methods were used to produce the sustainability scores for the concrete mixes: linear and geometric. linear aggregation is an additive operation (eq. 3), and is frequently the default method in mca, whereas geometric aggregation is a multiplicative operation (eq. 4). sin = n! i=1 (wi × scm ii) (3) sgeo = n" i=1 (scm ii) wi (4) where s: sustainability score, n: total number of indicators, scm ii: sustainability indicator i, and wi: weighting applied to indicator i. the choice of aggregation method represents another perspective on sustainability in the evaluation process. linear aggregation allows for full compensability between indicators; that is, an increase in one performance (such as co2 emissions) can be balanced by an equivalent decrease in another performance (such as cost). geometric aggregation, however, is less compensatory [12], meaning that trade-offs between indicators have a more pronouned effect on the sustainability score. the impact of compensability on the sustainability evaluation result may differ according to the conceptual framework, so these two 222 vol. 33/2022 evaluation of sustainable concrete �d��'luhfw�ordglqj��olqhdu�djjuhjdwlrq �e��'luhfw�ordglqj��jhrphwulf�djjuhjdwlrq �f��7kuhh�slooduv��olqhdu�djjuhjdwlrq �g��7kuhh�slooduv��jhrphwulf�djjuhjdwlrq �h��6'*v��olqhdu�djjuhjdwlrq �i��6'*v��jhrphwulf�djjuhjdwlrq� ���� ���� ���� ���� ���� ���� � �� �� �� �� �� �� �� 23&�1* 23&�5* )$�1* )$�5* %)6�1* %)6�5* 6 xv wd lq de lol w\ �v fr uh ���� ���� ���� ���� ���� ���� � �� �� �� �� �� �� �� 23&�1* 23&�5* )$�1* )$�5* %)6�1* %)6�5* 6 xv wd lq de lol w\ �v fr uh ���� ���� ���� ���� ���� ���� � �� �� �� �� �� �� �� 23&�1* 23&�5* )$�1* )$�5* %)6�1* %)6�5* 6 xv wd lq de lol w\ �v fr uh ���� ���� ���� ���� ���� ���� � �� �� �� �� �� �� �� 23&�1* 23&�5* )$�1* )$�5* %)6�1* %)6�5* 6 xv wd lq de lol w\ �v fr uh ���� ���� ���� ���� ���� ���� � �� �� �� �� �� �� �� 23&�1* 23&�5* )$�1* )$�5* %)6�1* %)6�5* 6 xv wd lq de lol w\ �v fr uh ���� ���� ���� ���� ���� ���� � �� �� �� �� �� �� �� 23&�1* 23&�5* )$�1* )$�5* %)6�1* %)6�5* 6 xv wd lq de lol w\ �v fr uh figure 3. sustainability scores by analytical scenario. aggregation methods were included in the mca process. 3. results and discussion 3.1. sustainability scores figure 3 shows the sustainability scores for the concrete mix alternatives by analytical scenario. for direct loading, the two bfs mixes exhibited the best sustainability performance, with bfs-ng slightly higher for linear aggregation, and bfs-rg slightly higher for geometric aggregation. in the case of the three pillars, fa-rg possessed the highest sustainability score regardless of aggregation method, followed by bfs-rg. finally, for the sdg concept, bfs-ng demonstrated the best sustainability for both aggregation methods, but the gap between the top four mixes was relatively small. overall, the four concrete mixes containing mineral admixtures consistently exhibited higher sustainability scores than the two opc mixes across all analytical scenarios. these mixes tended to have lower raw values for environment-related indicators compared to opc; and, since 12 indicators loaded to the environment, the effect of reduced environmental impacts was amplified through the mca process. while considering the limations of this analysis, these results nonetheless support the general perception that the application of alternative cementitious materials to concrete construction is a critical technological solution for improving sustainability in the concrete industry. 3.2. statistical examination it remains unclear which concrete mix is actually the most sustainable among the alternatives, as the results of each analytical scenario may be considered as equally valid due to the lack of an established stan223 michael henry, joel opon acta polytechnica ctu proceedings mix sustainability scores ranksmin. mean max. std. dev. variance mean variance opc-ng 34.3 37.0 38.5 1.6 2.2 6 3 opc-rg 35.8 38.1 42.1 2.6 5.8 5 4 fa-ng 53.4 55.0 56.7 1.3 1.4 4 2 fa-rg 54.6 56.8 61.2 3.1 8.0 2 5 bfs-ng 50.2 55.5 58.2 3.7 11.4 3 6 bfs-rg 56.3 57.6 58.4 1.0 0.8 1 1 table 4. summary statistics of the concrete mixes sustainability performance. analytical scenario (a) (b) (c) (d) (e) (f) (a) direct loading, linear aggregation 1.00 (b) direct loading, geometric aggregation 1.00* 1.00 (c) three pillars, linear aggregation 0.89 0.89 1.00 (d) three pillars, geometric aggregation 0.88 0.89 0.99 1.00 (e) sdgs, linear aggregation 0.99 0.99 0.86 0.86 1.00 (f) sdgs, geometric aggregation 0.99 0.99 0.86 0.86 1.00* 1.00 *actual coefficients were less than 1.00, but appear as 1.00 due to rounding table 5. correlation coefficients for sustainability scores between analytical scenarios. dard for concrete sustainability evaluation. to judge the sustainability under the uncertainty introduced by multiple concepts and aggregation methods, statistical properties of the sustainability score distributions for each material were examined (table 4). the mix with the highest mean sustainability score was bfs-rg; however, per figure 3, this mix was the most sustainable alternative in only one analytical scenario. on the other hand, bfs-ng was the most sustainable material in three scenarios, but its mean sustainability score was ranked third out of the six alternatives. similarly, farg was the most sustainable in two scenarios, but was ranked second by mean score. this result may be explained by the variance of the sustainability scores. the variance of bfs-rg is the lowest among all alternatives, suggesting it is relatively insensitive to the uncertainty introduced by different conceptual frameworks and aggregation methods, whereas the variance of bfs-ng is the highest, and thus the alternative most affected by the choice of concept and aggregation method (and their interactions). considering both the highest mean sustainability score and accompanying lowest variance, bfsrg may be judged as the most sustainable material when there is no consensus on the conceptual framework or aggregation method. finally, the pearson correlation coefficient matrix for the sustainability scores was calculated between analytical scenarios. it can be seen that all coefficients are equal to or greater than 0.86; in particular, the results of the linear and geometric aggregation scenarios were almost perfectly correlated for all three conceptual frameworks. this result suggests a very highly correlated set of sustainability scores regardless of the chosen conceptual framework or aggregation method, and despite the rankings of the mix alternatives shifting between scenarios. 4. conclusion this paper examined the effect of conceptualization on the sustainability evaluation of concrete using multicriteria analysis and a set of six concrete mix alternatives with varying binder contents and aggregate types. it was found that the most sustainable concrete mix varied according to conceptual framework and aggregation method, with three different mixes achieving the highest score among six scenarios. statistical examination of the distributions of the mixes’ sustainability scores revealed that the concrete mix with the highest mean sustainability score, bfs-rg, was only ranked first in one analytical scenario. however, the variance of this mix’s scores was the lowest among the alternatives, suggesting that its evaluation was comparatively less sensitive to methodological choices regarding conceptualization and aggregation. despite the different rankings for each analytical scenario, the sustainability scores were still very highly correlated between all scenarios, indicating that different methodological choices produced closely related results. it should be noted that the results reported here cannot simply be generalized to the greater concrete industry, as they are dependent on the sample and conditions of this specific analysis. however, it is expected that the demonstrated evaluation method will provide concrete industry stakeholders with an example of how to consider methodological uncertainties in sustainability evaluation to increase the robustness of their decision-making. 224 vol. 33/2022 evaluation of sustainable concrete references [1] k. sakai, t. noguchi. the sustainable use of concrete, crc press, pp. 188, 2017. 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[12] m. j. dobbie, d. dail. robustness and sensitivity of weighting and aggregation in constructing composite indices. ecological indicators 29:270-7, 2013. https://doi.org/10.1016/j.ecolind.2012.12.025. 225 https://doi.org/10.1016/j.jclepro.2019.01.220 https://doi.org/10.1016/j.conbuildmat.2014.01.039 https://doi.org/10.1016/j.autcon.2012.10.004 https://www.un.org/esa/sustdev/natlinfo/indicators/isdms2001/table_4.htm https://doi.org/10.1016/j.jclepro.2014.01.029 https://doi.org/10.1016/j.ecolind.2014.06.011 https://doi.org/10.1016/j.resconrec.2011.05.020 https://www.un.org/en/development/desa/population/migration/generalassembly/docs/globalcompact/a_res_70_1_e.pdf https://doi.org/10.1016/j.ecolind.2012.12.025 https://doi.org/10.14311/app.2022.33.0497 acta polytechnica ctu proceedings 33:497–503, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague a revival in façades: textile reinforced concrete panels are light, safe and aesthetically pleasing matthias roika, ∗, matthias tietzeb, alexander kahntc a halfen gmbh (part of leviat), liebigstr. 14, 40764 langenfeld, germany b technische universität dresden, dresden, germany c hochschule für technik, wirtschaft und kultur leipzig, leipzig, germany ∗ corresponding author: matthias.roik@halfen.de abstract. textile reinforced concrete was developed in recent years into a construction technique that has its benefits and advantages in applications where conventional types of reinforcement have their limits. the current, minimum slab thickness for steel-reinforced concrete façade slabs is 7.0 cm; this is due to the minimum, required concrete cover to ensure adequate corrosion protection. façade slab anchors for these slab thicknesses are building authority approved. as corrosion protection is not an issue for textile reinforced concrete, the minimum thickness for the concrete cover can be significantly reduced. the requirement for component thickness is now determined by the load-bearing capacity and by production-related boundary conditions. for practical building reasons, panel thicknesses of 3.0 cm have proven to be the best choice. compared to steel-reinforced façade panels, this is a weight and thickness reduction of almost 60%. thin concrete elements are of great interest in cases when the thickness or the weight of the panels is largely limited e.g. because of adjoining concrete elements in renovation or upgrade projects, retrofitting or improvements. compared to other building materials, concrete has characteristic advantages in building physics and fire protection properties, irrespective of the thickness. obviously, minimal thicknesses place extra demands on planning and construction. especially effects on concrete, punching, splitting and concrete breakout must be examined in experiments. this is an overview of calculation and test methods. results are provided to show the bearing behaviour of fixings in thin, textile reinforced concrete slabs. the design rules are explained and the results are illustrated. keywords: building innovation, carbon reinforced concrete façades, fixings, new materials, textile reinforced concrete. 1. introduction currently, more than half of the world’s population lives in cities. in 2050, this figure is expected to rise to two thirds. this will lead to a strong increase in the density of urban spaces, which in turn will lead to a high demand for façade constructions that are efficient in terms of raw materials and space. the general availability of raw materials, sufficient loadbearing capacity and durability, as well as the ease of production have made reinforced concrete the building material of the 20th and the beginning 21st century. in 2017, approximately 4.3 billion tonnes of cement, 28.6 billion tonnes of aggregates and 2.85 billion tonnes of water were used to produce concrete [1]. and this is the other side of the coin of the universal building material concrete: no other material is responsible for greater raw material extraction and higher co2 emissions. the huge quantities of concrete used worldwide are particularly problematic. if it were possible to achieve a leaner design through a new type of construction, considerable savings could be achieved carbon concrete offers this potential. reinforcing structures (grids) made of carbon fibers can already be used today to replace some steel reinforced concrete applications in new buildings, i.e. to achieve greater resource and energy efficiency [2]. in addition, it is possible to repair solid structures in need of renovation by means of reinforcing layers of carbon concrete, thereby increasing their service life. for architects and designers, carbon concrete and other types of textile-reinforced concrete offer great design freedom, as they can be used to produce any shape, format, surface structure and colour in high quality. furthermore, additional functions such as heating, lighting or building automation can be integrated into the components. due to the arrangement of the reinforcement according to the force distribution, components made of textile concrete also require less reinforcement than conventional, fibrereinforced concrete, in which cut short fibres made of glass, plastic or carbon fibre are introduced into the concrete mix in a non-directional manner. in germany, experiments with reinforcement made of technical textiles were already being carried out in the 1980s. the first joint research projects in dresden, chemnitz and aachen followed in the mid-1990s. they formed the basis for two dfg-funded collab497 https://doi.org/10.14311/app.2022.33.0497 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en m. roik, m. tietze, a. kahnt acta polytechnica ctu proceedings figure 1. principle cross-section curtain wall façade. orative research centres in dresden [3] and aachen [4], in which basic research in the field of non-metallic reinforcements was advanced between 1999 and 2011. in what is currently the largest german construction research project c3 carbon concrete composite more than 150 partners from science and industry are working on the development and market launch of carbon concrete. carbon concrete will change the construction world, the technology is already there and is gradually being used in more and more areas of construction due to its advantages. the first applications have been realized in the field of façades, in addition to the repair of reinforced concrete structures. through the further development of the components concrete matrix and carbon reinforcement of the composite material, further fields of application can now be opened up. carbon concrete offers the possibility to rethink old approaches in the building industry. 2. carbon reinforced concrete in reinforced concrete, the concrete primarily absorbs the compressive forces occurring in the component and the reinforcing steel compensates for tensile stresses. in addition, the concrete cover protects the steel against corrosion, corrosion-promoting media (e.g. de-icing salts) and fire. depending on the paving situation, the concrete cover course must be up to 5 cm thick. in the case of carbon concrete, car� figure 2. carbon reinf. concrete façade white concrete curtain wall. bon, as the equivalent of steel reinforcement, ensures the tensile strength of the composite material. the principles and design bases are very similar to those of reinforced concrete. since carbon is about six times more load-bearing (3000 instead of 500 n/mm2 tensile strength) and four times lighter (1.8 instead of 7.8 g/cm3) than steel and also does not corrode, this reinforcement allows a much freer design. since the corrosion protection for the reinforcement is no longer required, concrete slabs for façade cladding, for example, can be produced with a thickness of 2 cm instead of the previous minimum of 7 to 10 cm [5]. even thinner layers are sufficient for reinforcing existing reinforced concrete components. currently, filigree mat-like structures made of one or more layers of multifilament yarns (rovings) are used for reinforcement. rovings in turn consist of several thousand individual filaments (fibres). since the fabrics are relatively soft, they can be used to produce almost any shape. however, for entire structures, rod-shaped reinforcements will be required in the future. various research institutes and industrial companies in germany are currently working on the development of suitable carbon rods. in addition to carbon, fibers made of ar glass are becoming increasingly important for multifunctional applications. these are somewhat less load-bearing (1500 n/mm2) and durable than carbon fibers, but are significantly less expensive. furthermore, unlike carbon, they do not conduct electricity, but are good light guides. up to now, normalstrength or high-strength fine concrete with a grain size of 1 to 8 mm was used as concrete depending on the component and application. this is also a difference to steel reinforced concrete, where grains up to 32 mm in diameter are common. this concrete matrix is particularly suitable for thin building components such as façade panels [6]. 3. carbon concrete façades for façade construction, carbon-reinforced concrete elements offer the advantage that they have very uniform, fine-pored and sharp-edged surfaces and can be produced much slimmer and lighter than steel re498 vol. 33/2022 textile reinforced concrete panels figure 3. storey-high façade element with new fixing system for carbon reinforced concrete. inforced concrete elements. the reduction in weight leads to greater economic efficiency in production and transport and reduces environmental pollution. the elements are mainly used for curtain wall rear ventilated façades (figure 1). furthermore, the panels are suitable for the production of particularly slim sandwich façades. the first office and commercial building with a curtain-wall ventilated façade made of carbon concrete was completed in neumarkt in 2016. the 2.88 * 4.48 m large façade elements here are only 30 mm thick. a reinforced concrete facing layer would have been around three times as thick at this point. the white concrete façade elements on a modular single-family house in leipzig are also only 30 mm thick and reinforced with a carbon fiber fabric (fig. 2). the panels are up to 0.75 × 3.00 m in size and are held in place by a clasp system. an office building in essen shows the possibilities of three-dimensional surface design. here, too, a curtain-wall, rear-ventilated façade was realized, whereby the concrete slabs are anthracitecolored throughout and provided with a binary code with the help of a formwork matrix. a fair-faced concrete façade element only 11 cm thick with passive house insulation standard (0.15 w/m2k) was developed within the framework of the research project "vakutex vacuum-insulated façade elements made of textile concrete" in leipzig [7]. the non-load-bearing sandwich element has similar physical properties as a four times thicker reinforced concrete element. this results in floor space gains of up to 15 % for the same building cubature. for the implementation of the first projects in the façade area, approvals were required in individual cases. in order to make this easier, general building authority approvals were obtained, which allow the approved component to be used throughout germany. façade elements have also been used in non-european countries. the pylons of the new borfigure 4. construction of the façade in dormettingen (ger). porus bridge yavuz-sultan-selim were clad with textile and carbon concrete façade elements at a height of 326m. for the further practical suitability of the technology, a façade anchor system was developed which is particularly suitable for large-format thin (min. d = 3 cm) concrete elements. with this system, concrete façade panels and ventilated façades made of textile and carbon reinforced concrete can be easily installed and adjusted. this technology has already been implemented in a practical project, the schiefererlebnis in dormettingen (germany). (figure 3 and 4) this shows that carbon-reinforced concrete technology has arrived in practice. 4. fixtures in concrete after the bearing construction is erected, the façade panels will be fixed. inserts are necessary for handling and support to provide a safe connection between the bearing construction and the panels. the general approach to calculate anchors in concrete is the cc (concrete capacity) method [8, 9]. the mechanisms of load-transfer from a steel anchor into the concrete are either mechanical interlock, friction or bond. anchors can be cast-in-place or post installed. usually, cast-in anchors are fixed by mechanical interlock whereas post installed anchors, such as dowels placed in a drilled hole introduce a load by friction or bond. raising the load and assuming the anchor material is strong enough to carry the tensile load, the fixing fails because the tensile capacity of the concrete is exceeded. this failure mode has a shape of a cone with a slope of approximately 35◦ from the horizontal. the concrete cone failure load is proportional to h 1,5 ef , where hef is the effective anchorage depth, the depth where the anchor introduces the load into the concrete. the concrete capacity method is based on calculation of this failure cone. the cone capacity can be influenced by geometry of the concrete member, by reinforcement or by additional cracks. 499 m. roik, m. tietze, a. kahnt acta polytechnica ctu proceedings figure 5. stress distribution in uncracked concrete. figure 6. stress distribution in cracked concrete due to splitting. 5. experimental approach tests have been carried out to investigate the support behaviour of fixings in thin concrete slabs: calculations according to the cc-method set the minimum depth of anchorage at hef ≥ 25 mm. this minimum anchor depth is not feasible for applications with carbon reinforced concrete and the calculation method should be reviewed for the use case of thin slabs. tests were carried out to extend the effective range of the anchors for short depths. these tests were carried out on thin concrete panels, with and without reinforcement. 5.0.1. punctual fixings additionally to [10], tests with headed bolts installed in 30 mm thick concrete plates were carried out to investigate the bearing behaviour of anchors installed in thin plates. bolts were installed at the maximum feasible anchorage depth and the shaft diameters and different head sizes varied. all bolts were tested in tension. first results were published in [11], additional test series followed. unexpectedly large scatterings of the test results were observed in all test series. these results were not sufficiently reliable due to a coefficient of variation well in excess of 20 %. despite the observed tendency for the resistance of concrete fracture bodies to depend directly on the lateral area, it was not possible to derive a practically applicable mathematical rule for this dependence. the cause of these large scatterings can be attributed to the fracture mechanism itself. the order of crack initiation was assessed, and it was found that the initial surface cracks occurred during the test. the final failure of figure 7. supporting mechanism of a head bolt anchorage due to shear load (eligehausen, mallée & silva 2006) figure 8. load capacity depending on the relation between axial / lateral load, (dafstb 2008). the anchor is almost always due to splitting rather than the expected failure of the concrete. the failure mechanism of splitting is shown in figure 6. the compressive load introduced into the concrete by the anchor head results in tensile stresses near the surface of the concrete member, as shown in figure 5. if initial cracking occurs near the anchor due to bending stress in the concrete, further splitting stresses will open the cracks and reduce their resistance. therefore, fixtures which are sensitive to splitting failures will have a lower resistance value and a higher variability of test results. if a fixture is sensitive to splitting, a corresponding coefficient should be introduced [12]: ψh, sp = ! h 2 · hef "2/3 ≤ 1.0 (1) this factor reduces the basic resistance nrk,c in case when h < 2 hef . when additional shear forces are applied to the concrete, the tensile load component of the headed anchor increases (see figure 7). this additional load component can lead to so-called prying failure, which also leads to a reduction in the overall tensile resistance. this effect is investigated in figure 8 [10]. loads were applied to the anchors at large distances and at different angles towards the edge of the panel. the resulting crack diagrams in the report show additional 500 vol. 33/2022 textile reinforced concrete panels figure 9. tests with additional reinforcement. the resistance of the unreinforced concrete forms the reference value. it can be seen from the pictures that the load-bearing capacity is significantly influenced by the crack development. the inserted carbon meshes prevented cracks and a concrete breakout cone could form. figure 10. fixing system for façade panels. splitting and fracture of the concrete cone, in addition to the overlapping of the cracks due to prying failure. this may lead to a further increase in scattering (see figure 8). finally, it can be said that timed fixings are not very suitable for applying large loads to thin fibrereinforced panels, especially when the loads contain a large shear component or when h < 2 − hef . depending on the size of the panel, the resistance of the fixings may not be sufficient to support the façade panel. 6. methods to reduce the influence of splitting to increase the load-bearing capacity of an anchorage and to reduce the scattering of the test results, various measures can be taken. two measures are described in more detail below, on the one hand the use of reinforcement and on the other hand the distribution of the load to be introduced over several figure 11. test setup for shear/tension. load-bearing components. 6.1. use of additional reinforcement additional reinforcement can be arranged in several ways. in addition, different types of reinforcement were considered in tests, namely bar-shaped reinforcement and reinforcement mats. in the case of barshaped reinforcement, it must have its own corrosion protection due to the low concrete cover, i.e. it should consist of e.g. stainless steel or carbon. the same reinforcement was used as mat-shaped additional reinforcement, which is also used as slab reinforcement. the results of the tests with additional reinforcement are shown in figure 9 6.2. multi directional fixings special anchoring systems with suitable insert elements are required to fix concrete facades. to prevent eccentricity of the supporting structure, it is advantageous that the insert can be loaded in different directions. it is convenient to equip this type of insert element with several components to introduce loads independently into the concrete. an example of this is shown in figure 10. headed anchors carry the loads 501 m. roik, m. tietze, a. kahnt acta polytechnica ctu proceedings figure 12. test results vs. calculation results of multidirectional fixing. perpendicular to the surface of the facade, i.e. tensile loads, while reinforcement bars, for example, support the shear direction. these inserts were subjected to tensile tests at different angles. the test set-up is shown in figure 11. the concrete panels were fixed to a special test rig and loaded at different angles. the insert system was designed to accommodate loading angles of between 25◦ and 45◦. this range was used for the tests. it was found that even when relatively large loads were introduced into the concrete, the loads did not depend significantly on the loading angle. the final fractures were of sufficient size to achieve a noticeably higher resistance than the time-fixed ones; for values below 10%, the scatter was well within acceptable limits. these findings demonstrate that multidirectional fixings have significantly better loadbearing properties than the fixings. these inserts were subjected to tensile tests at different angles. the test set-up is shown in figure 11. the concrete panels were fixed to a special test rig and loaded at different angles. the insert system was designed to accommodate loading angles of between 25◦ to 45◦. this range was used for the tests. it was found that even when relatively large loads were introduced into the concrete, the loads did not depend significantly on the loading angle. the final fractures were of sufficient size to achieve a noticeably higher resistance than the time-fixed ones; for values below 10%, the scatter was well within acceptable limits. these findings demonstrate that multidirectional fixings have significantly better loadbearing properties than the fixings. the results of the test are shown in figure 12. to describe this interaction, a computational model has been developed according to [8, 9]. the test results are presented together with the relevant safety levels according to [13, 14]; the difference between the safety levels in [13, 14] is based on the difference in confidence levels (the confidence factor in [13] is 75% and in [14] it is 90%). in the case of reinforced concrete structures, [13] is the restraining force and in the case of fixed structures, [14] is the restraining force. tests have shown sufficient safety and a good correlation with the calculated values. the inserts have been specially developed for thin carbon-reinforced concrete panels and have been approved by the general building authority. for the horizontal load component due to wind loads, a spacer such as a compression element is usually installed. this compression element can take the horizontal load without affecting the load-bearing equipment. the thinner plates are more sensitive to wind loads, as they carry less dead load than thicker concrete façade panels. depending on the shape of the slab, the weight of the slab and the expected wind loads, horizontal anchors (e.g. compression bolts or restraint anchors) may need to be additionally fixed for wind suction. in stacked, suspended façade slabs, the lower spacer bolts are replaced by dowels for aligning the elements. these elements are also included in the general building approval. 7. numerical investigations numerical investigations of the effects observed in the experiment are in progress. based on the undisturbed concrete failure cone, i.e. concrete breakout without edge and thickness influence, the resulting resistances are currently to be determined, which occur if the influencing factors are guided to the limits of applicability. 502 vol. 33/2022 textile reinforced concrete panels as can already be seen from the test results, the influence of the splitting coefficient ψh,sp, for example, is considerably greater than stated in current publications. the influence of the reinforcement is currently being investigated. 8. summary / outlook the construction of carbon reinforced concrete can be described as a mega-trend, as this technology will be used in many areas of the construction industry. façade construction in particular has been one of the first areas of application to show the advantages of carbon reinforced concrete. above all, the savings of resources is an essential aspect which, in addition to the many technically advantageous properties, has benefits for society as a whole. of course, not all design methods known from reinforced concrete can be transferred to carbon concrete. therefore, new design methods are being developed at the universities in dresden and aachen (germany). in addition, the current technical specifications for fastening techniques developed for conventional steel-reinforced structures must be adapted to the specific boundary conditions of carbon concrete construction. for this purpose it is necessary to adapt the relevant parameters to the specific properties of the fabrics (reinforcement), which is usually done by tests. although the effective anchorage depth and the edge distance of an anchor according to the formula should have the greatest influence on the load bearing behaviour, the influence of cracking of the concrete dominates. in further investigations by means of tests and numerical analysis further special features of carbon reinforced concrete are to be analysed, such as the influence of the special bond behaviour or the reduced shear resistance compared to steel reinforcement, in order to derive efficient systems for practical use. acknowledgements we would like to thank the german federal ministry of education and research for its support of carbon concrete research within the framework of the twenty20 program, and our corporate partners for their successful cooperation. references [1] mineral commodity summaries 2019, 2019. https://doi.org/10.3133/70202434. [2] a. kahnt, s. kirmse, k. holschemacher. global challenges for the construction industry. investicii, stroitel’stvo i nedvizhimost kak materialnyi bazis modernizacyi i innovacionnovo razvitya ekonomiki, p. 55-59, 2014. [3] sfb 528, http://www.sfb528.tu-dresden.de. [4] sfb 532, http://www.sfb532.rwth-aachen.de. [5] scheerer, s.; schladitz, f.; curbach, m.: textile reinforced concrete from the idea to a high performance material. in: brameshuber, w. (hg.): proceedings of the ferro-11 and 3rd iccarbon reinforced concrete (pro 98), aachen, bagneux, 15-33 2015,. [6] f. schladitz, m. tietze, m. lieboldt, et al. carbon and concrete the future of construction? iabse symposium report, iabse conference guangzhou 2016: bridges and structures sustainability-seeking intelligent solutions, pp. 733-743, 2016. https://doi.org/10.2749/222137816819259167. [7] s. kirmse, a. kahnt, r. kraft, et al. energy façades made of textile reinforced concrete. cpi concrete plant international. issue (3):158-171, 2017. [8] cen tc250/sc2/wg2. design of fastenings for use in concrete, prcen/ts 1992-4-1:2009, 2009. [9] fib. fédération international du béton. design of fastenings in concrete guide to good practice parts 1 to 5, lausanne, 2011. [10] dafstb. concentrated load application into thin-walled components made of textile reinforced concrete, dresden, 2008. [11] r. matthias. new equipment for new technologie: fixing systems for thin, textile reinforced concrete facades, icsa 2016, guimaraes, portugal, 2016. [12] r. eligehausen, r. mallée, j. f. silva. anchorage in concrete construction, ernst und sohn, 2006. [13] din en 1990:2010-12 eurocode: basis of structural design; german version en 1990:2002 + a1:2005 + a1:2005/ac:2010, 2010. [14] etag 001. eota: metal anchors for use in concrete, 2013. https: //www.eota.eu/sites/default/files/uploads/etag s/etag-001-annex-a-amended-2013-04-08.pdf. 503 https://doi.org/10.3133/70202434 http://www.sfb528.tu-dresden.de http://www.sfb532.rwth-aachen.de https://doi.org/10.2749/222137816819259167 https://www.eota.eu/sites/default/files/uploads/etags/etag-001-annex-a-amended-2013-04-08.pdf acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0183 acta polytechnica ctu proceedings 38:183–189, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague digitalising cities: a methodology to map evaluation requirements into robust and feasible data collection approaches josé l. hernándeza, ana quijanoa, ∗, pierre noailleb, mikko virtanenc, rubén garcíaa a cartif technology centre, energy division, parque tecnológico boecillo 205, 47151, boecillo, spain b cerema, direction territoriale ouest, man, rue rené viviani, 44000, nantes, france c vtt technical research centre of finland ltd., p.o. box 1000, 02044, espoo, finland ∗ corresponding author: anaqui@cartif.es abstract. the sustainable development of cities relies on the implementation of multi-sectoral actions towards carbon neutrality, reducing the air pollutants emissions. the actions’ decision-making process for cities transformation should be supported by lessons learnt from previous interventions and kpis (key performance indicators). to do so, gathering real data becomes pivotal, complementing simulation tools (currently used), solving the inherent uncertainties due to assumptions. data collection methodologies are then necessary, being the main driver for digital cities and providing better mechanisms for informed decision-making. most of the cities still operate in silos and do not always implement the strategic plans supported with a digitalization of the municipal processes. within this perspective, this paper presents a methodology to support cities in the preparation of monitoring programmes to collect real data in a robust and feasible manner. taking the kpis and the smart cities urban strategies into account, this paper concludes with some lessons learnt within cities to deploy monitoring approaches. from the city challenges to the review of the plans, all the process is driven by real data and kpis. the methodology has been applied in the mysmartlife project (grant agreement #731297) and deployed into the cities of nantes (france), hamburg (germany) and helsinki (finland). keywords: digitalisation, sustainable cities, kpis-based assessment, monitoring, decision-making, data. 1. introduction cities population is growing and nowadays, more than two thirds of the european population (the 70.9 % [1]) is living in urban areas [2], where the 29.1 % of people live in rural areas [1] with expectations to increase the statistic in the next years. this is partially because of the perception of cities as centres of economic growth that provide opportunities for study, innovation and employment [2]. additionally, cities are responsible for more than 60 % of greenhouse gas emissions, a figure that will increase with the population growth. for these reasons, there is a need for a transition, not only from current cities into smart cities, but also, as stated in the eu green deal and the horizon europe climate-neutral and smart cities mission, towards a more ambitious target that aims at the creation of climate-neutral cities. these are cities that optimize the resources, enhance the citizens’ participation, increase the satisfaction of living, as well as quality of live and sustainability [3]. however, the major challenge lies in how to assess the level of smart cities and the achievement of the sustainable goals. many cities implement subjective criteria to evaluate themselves; hence, objective methodologies and indicators are necessary, with certain level of uniformity [4], so as to be able to determine the strengths and weaknesses of the municipality. these methodologies and indicators should be driven by real data, but cities fail in the definition of proper strategies to collect real data [5, 6]. on the one hand, they lack of expertise in the application of technologies like iot [5], being usually supported by external parties. on the other hand, cities are split into verticals (e.g. energy or mobility), without relationships between councillorships [6]. cities are currently working on new urban transformation or regeneration models to transform current practices towards a smart city [4]. these require the dependency between various perspectives of the city, e.g. impacts on the energy demand due to the electrification of the mobility. then, smart cities of the future are a combination of multiple pillars such as energy, mobility, nature, economy or water management, among others, but always with the citizens at the core. all of these are supported by the integration of the information and communication technologies (ict) [7]. within this perspective, mysmartlife project [8] aims at the transformation of the cities of nantes, hamburg and helsinki into more environmentally friendly, as well as more liveable cities, by reducing greenhouse gas emissions, supported by the digital183 https://doi.org/10.14311/app.2022.38.0183 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en j. l. hernández, a. quijano, p. noaille et al. acta polytechnica ctu proceedings isation process. more than 150 actions have been implemented across the three cities following the concept of “inclusive cities”, where citizens play a key role through the smart people concept. furthermore, smart economy is the driver for the urban transformation, attracting talent and providing goods and offering services accompanied by innovative business models to satisfy the requirements and creates a solid business case. the final impacts of these actions need to be calculated to extract conclusions. under this objective, mysmartlife has defined a holistic evaluation procedure [9], where key performance indicators (kpis) are set to facilitate the impact assessment. the mysmartlife framework merges multiple pillars (energy, mobility, social, economy, ict and governance) of the city to provide a holistic analysis of the smart city. the implementation and deployment of this procedure relies on data, which is presented as one of the main challenges [5]. to overcome it, this paper presents a methodology that has been applied in nantes, hamburg and helsinki to define robust and feasible monitoring programmes with the aim of collecting real data that support the calculation of the kpis. it should be noted that the paper is focused on the digitalisation of the cities under the pillars of energy and mobility, whose assessment is based on real data, leaving out of the scope the analysis of qualitative data for other verticals. this is an iterative process for the implementation of smart city strategic plans, which can be adapted in each step to support better informed evaluation and, thus, decision-making processes. as results, this paper extracts examples and lessons learnt of the cities of nantes and helsinki. the rest of the paper is organised as follows. section 2 provides a background of current practices. section 3 presents the methodology that has been proposed within mysmartlife to approach the urban plans and data-driven evaluation. finally, section 4 shows the examples for nantes and helsinki. section 5 extracts the conclusions and lessons learnt of the proposed methodology. 2. background and current practices in smart cities as it was introduced before, the smart city concept merges multiple pillars, such as energy and environment, sustainable mobility, smart people, smart economy and transparent government, among others. all of them are supported and enabled by the implementation of ict solutions for digitalisation and data management. however, there is no consensus about standard procedures for assessing the achievement of the city goals within these pillars [4]. for instance, the authors in [10] studied 16 evaluation frameworks, including more than 950 indicators, concluding the lack of sustainability assessment. also, the ranking of indicators is highly conditioned by the data availability [4], which complicates the application of data-driven approaches for smart cities evaluation. open initiatives, like the one presented in [11], offer new frameworks under which impact and sustainable evaluation can be carried out. the result of [11] is mainly theoretical, but fosters the open data accessibility to calculate indicators. nevertheless, the reality of a city is different. the challenge is not the exposition of data, but the collection from field, requiring the definition of robust monitoring programmes to be able to ingest data before being openly shared. this is of special importance when private companies are involved. apart from data availability, it should be considered the heterogeneity of cities and its complexity in the governance aspects, being one of the causes the involvement of different actors (i.e. the councillors of the municipality, urban planners, citizens, private companies, investors . . . ). the convergence is pursued by the methodology that is presented in [12]. within this analysis, building and district renovation, urban mobility, icts and non-technical actions are considered under the umbrella of a data collection approach. the authors in [13] presented a survey, which includes smart nature, smart living, smart mobility, smart governance, smart people and smart economy, without a clear data gathering approach to evaluate the aforementioned verticals, similar to the case in [11]. the examples of current initiatives for city impacts assessment converge in the need for kpi-driven evaluation framework. however, no one considers the mature level of the city in terms of digitalization or, in other words, data to calculate the indicators. while some cities are highly digitalised, with a wide deployment of sensors across the city, others that are still using analogue data. this paper then presents a methodology to support cities in the digital transformation to overcome the barriers between data collection and kpi calculation. as far as it is known by the authors, there are not researches regarding this line of work. literature is focused on the analysis of evaluation frameworks, such as [10], without caring about the procedure to assure real data to apply them. 3. mysmartlife methodology: translating kpis into monitoring this section provides an overview of the methodology that has been applied within mysmartlife project, which is depicted in figure 1. it consists of four stages, described below, which are cyclical and iterative. cities need to be continuously transforming and adapting to changes; therefore, this methodology allows dynamic and continuous evaluation. the starting point of the approach is the availability of any assessment framework with well-established kpis, which will be the reference for the city. in this case, mysmartlife counts on its own framework [9]. the 184 vol. 38/2022 digitalising cities: a methodology to map evaluation requirements . . . figure 1. roadmap for a data-driven sustainable environment evaluation. four stages of the proposed methodology are explained in the next bullets: (1.) analysis of the city. being a cyclic and iterative methodology, this phase represents at the same time the initial and final stage of each one of the iterations. it aims at quantitatively determining the city challenges (i.e. by using the kpis from the evaluation framework). the outcome of this stage is to obtain the city audit. in the first iteration, this analysis is made at macro level (i.e. city as a whole), meanwhile, in the rest of iterations it is micro level, i.e. evaluating the actions that have been implemented in the smart city plans. this quantitative analysis, at the end of the iteration, allows planning the next steps by extracting the lessons learnt and knowledge from real data. (2.) smart city plans. from the previous analysis of the city, where the challenges and needs are gathered, the actions and interventions for the sustainable transformation are determined in this stage. these actions and interventions contain clear objectives and targets to be achieved, e.g. reducing of energy demand in 30 %. the selection of the most suitable actions is based on benchmarking through estimated values from the indicators, such as the case of the morgenstadt initiative [14]. it is normal that the estimations were performed with simulation tools, which make some assumptions that are not representing the real situation. within the methodology being described in this paper, real data support the estimation thanks to the digitalisation strategy and icts. the main advantage or benefit is the capability of selecting actions under quantitative evidences, instead of experiences, which are usually biased by subjective perceptions. (3.) kpi selection for actions’ evaluation. within this stage, the most suitable kpis from the framework [9] are extracted according to the objectives of the actions. the framework [9] considers five axes: smart environment (that covers energy and greenhouse emissions), smart mobility, smart people, smart economy and smart governance, as well as icts in the form of urban platforms for the digitalisation of the city. more than 150 indicators are available. (4.) monitoring stage. this is the main challenge of the cities. current decisions are based on simulation tools that make some assumptions, introducing errors. data are crucial to calibrate the models or make data-driven decision-making. this phase establishes the requirements about what to measure according to the kpis and where to measure (location of sensors), as well as how (e.g. iot pollution sensors or heat meters), when (frequency) and who (responsible). many of the studies do not include this crucial stage, which is in charge of the data collection for assuring data-driven strategic plans for sustainable transformation. this methodology is complemented with the urban data platforms, with the goal of storing data in a persistent way that is defined by the monitoring programmes [15]. the urban data platform makes data available for a later analysis and data-driven decision making. the monitoring programmes are a theoretical definition of the datasets and data-points that are required from the kpis perspective, but its deployment is not covered by the proposed methodology. in this case, the cities, together with their local innovation ecosystems and through the deployment of sensors and the integration of the related datasets 185 j. l. hernández, a. quijano, p. noaille et al. acta polytechnica ctu proceedings figure 2. kpi-driven assessment framework. smart smart smart smart smart environment mobility people economy governance city (macro-) level 56 22 16 16 15 project (micro-) level 32 51 5 22 7 table 1. total number of kpis defined for each of the pillars. in the data platforms, are the responsible for managing the physical interventions. the main advantage of using the urban data platforms is the availability of data in open formats for further treatment and kpi calculation. the automation of the calculation algorithms and decision-making dashboards is out of scope of this paper, although it is recommended for the analysis [15]. by following the proposed methodology, one of the main benefits is to keep aligned the objectives and challenges of the city, which are driven by quantitative evidences coming from data. data collection and knowledge extraction support the analysis of the actions impact and lessons learnt to enhance sustainable development based on better-informed decisionmaking. data are the complement for the simulation tools in order to calibrate them and extract real performance metrics. in this way, the assumptions taken when creating the simulation models and uncertainties are demonstrated or solved. all in all, this methodology presents a holistic approach for the cities to continuously improve the sustainable transition. on one hand, by updating for each cycle the indicators of sustainable development; on the other hand, by aligning objectives, kpis and real data. 3.1. kpi-driven assessment framework the methodology relies on the existing kpi-based evaluation framework. this section summarises the evaluation framework under which the methodology is supported. the concept is represented in figure 2 [9]. as it can be observed, there are two levels: city (providing the smart city vision) and project (the mysmartlife actions). the city level extrapolates the results of the actions to the whole city to determine the benefits under the assumption of implementing these interventions across the city. both levels cover the same six categories (named fields in the city level and pillars in the project level). although the aim of this paper is not the definition of the kpis, which are available on [9], table 1 summarises the kpis account for each pillar, excluding the digitalisation metrics. all are included within the overall framework of the, but only the selected kpis (as per step 3) will be evaluated in the iterative analysis of the city, being chosen considering the strategic plan objectives and actions. 3.2. mapping kpis into monitoring requirements the key for the success of the application of this datadriven procedure is the capability for collecting real data from the different city resources. some cities fail in the collection of data so as to achieve a specific objective and they just acquire data [16] that is not exploited. then, a comprehensive and robust data collection approach needs to be applied. for that end, a common template has been provided to support cities in the mapping of the objectives, kpis and monitoring requirements. this template is focused on energy and mobility, which require the deployment of physical devices, while social, economic, governance and ict (digitalisation) aspects do not require such monitoring schemas, although definition of data collection methods like surveys and questionnaires are also used. 186 vol. 38/2022 digitalising cities: a methodology to map evaluation requirements . . . figure 3. monitoring template schema for energy monitoring in cities. figure 3 [17] represents the schematic view for the energy monitoring. on right side, it shows the thermal and electrical generation systems that could be the grid, renewables, district heating or any other supplying energy. in the middle of the schema, there is the assessment boundary, where the interventions apply, which is pivotal to determine the area of intervention [9]. in the example, the boundary covers both the buildings (e.g. retrofitting) and energy distributions elements (e.g. heat exchangers). this virtual boundary is like a geographical or physical one that splits the systems/interventions to be analysed (those that are part of the smart city plan for sustainability) from the existing elements (e.g. the electricity grid). this boundary is flexible and can be completely adapted to the strategic plan of the city. finally, on the left side, it is shown the demand side, e.g. the buildings that are retrofitted or new construction, which are the final “users” of the energy. also, the monitoring template allows cities to identify the energy flows and the connection between elements. while a grey arrow points the thermal energy flows, a black arrow identifies the electricity flows. in this way, the elements of the physical connections are easily identified and where to measure by placing directly in the schema the icon representing the measurement sensor (e.g. wattmeter). each of these sensors represents one measurable parameter of the kpis. it should be noted that this schematic definition is flexible enough to be adapted to the needs of various cities. it aims at following the same concept of boundaries, energy flows and so on. nantes and helsinki have already adapted this concept to their specific contexts. moreover, it should be remarked that zoom-in is possible to specific interventions, such as district heating details. 4. examples of application in the cities this section aims to show direct examples of the application of this methodology to define monitoring schemas in nantes and helsinki lighthouse cities within the framework of mysmartlife project. figure 4. monitoring programme for district heating in nantes. 4.1. nantes case nantes is a city committed with the sustainable transformation for climate change adaptation. for that end, sustainable energy action plan (seap) and sustainable urban mobility plans (sump) were developed as initial strategic plans based on macro-indicators. these plans have driven the calculation of the city audits and baseline to establish the targets of the actions defined in mysmartlife [18]. the main interventions defined in the smart city plans are the renovation of the building stock, renewable district heating, electro-mobility, citizens’ involvement and engagement. under this perspective, the evaluation framework defined during the first stage of the project was adapted to these plans by selecting the significative kpis for the nantes evaluation. following these kpis, the monitoring programme was defined as depicted in figure 4 [17]. it reflects the flexibility of the template, where a zoom is made to the district heating, which has to assure 80 % of renewable contribution, according to the smart city plans. in this sense, the contribution for each source is measured, both renewable and non-renewable to obtain the final ratio. with the results, at the end of the project, the fulfilment of the plan will be evaluated (by means of the kpis) to be reviewed. 4.2. helsinki case helsinki started from initial seap and sump [19], whose baseline and targets were established based on the selected actions. kpis adapt the evaluation 187 j. l. hernández, a. quijano, p. noaille et al. acta polytechnica ctu proceedings figure 5. monitoring example for the viiki building in helsinki. framework to the necessities of these strategic plans. throughout the process, existing data sources for the kpi calculation were mapped and possible blind spots in terms of monitoring identified and, thus, obtaining the monitoring needs for the impact calculation. by applying this data-driven methodology, the monitoring schema for high-performance office building (equipped with a rather sophisticated energy system consisting of both thermal and electrical grid connections, renewable energy generation and storage) is detailed in figure 5 [17]. the template is able to encase the core needs for monitoring in terms of final calculation of the kpis. 5. conclusions and lessons learnt in this paper it has been described a methodology to support cities in the digitalisation process to deploy impact assessment frameworks when evaluating sustainable plans for climate change adaptation. the methodology provides guidance in the data collection strategies to supply quantitative evidences in the decision-making processes, while qualitative analysis is out of the scope of this paper. the presented mysmartlife methodology benefits cities in the capabilities to deploy monitoring programmes focused on well-established kpis and evaluation frameworks. the iterative nature of the methodology grants an adaptative capacity in the smart city strategic plans for sustainable transition. the methodology has been deployed in the cities of nantes, hamburg and helsinki, although this paper presented the cases of nantes and helsinki. the main lesson learnt or conclusion, extracted from the cities, is the support provided by a visual tool to determine the monitoring programmes. these help to calculate audits and baselines and apply better-informed decision-making processes. in contrast, another lesson learnt is the complexity in the management of the completion of the schemas because they require the interaction among municipality stakeholders, which is not always easy. acknowledgements the authors would like to thank, first, the ec for funding and supporting mysmartlife project under ga#731297. secondly, the authors would like to thank the rest of consortium of mysmartlife for the contributions in the definition of the framework and roadmap. references [1] eurostat. urban and rural living in the eu. 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[2022-02-06], https://www.mysmartlife.eu/fileadmin/user_ upload/deliverables/d4.1_baseline_report_of_ helsinki_demonstration_area.pdf. 189 https://doi.org/10.3390/environsciproc2021011021 https://doi.org/10.1016/j.cities.2016.09.009 https://doi.org/10.1016/j.cities.2020.102860 https://doi.org/10.3390/su13179683 https://doi.org/10.1109/seaa.2018.00084 https://www.morgenstadt.de/en.html https://doi.org/10.3390/s20082402 https://doi.org/10.3390/s17081834 https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d5.3_monitoring_programs_and_deployment_in_the_three_lighthouses_stamped.pdf https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d5.3_monitoring_programs_and_deployment_in_the_three_lighthouses_stamped.pdf https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d5.3_monitoring_programs_and_deployment_in_the_three_lighthouses_stamped.pdf https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d5.3_monitoring_programs_and_deployment_in_the_three_lighthouses_stamped.pdf https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d2.1_baseline_report_of_nantes_demonstration_area.pdf https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d2.1_baseline_report_of_nantes_demonstration_area.pdf https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d2.1_baseline_report_of_nantes_demonstration_area.pdf https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d4.1_baseline_report_of_helsinki_demonstration_area.pdf https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d4.1_baseline_report_of_helsinki_demonstration_area.pdf https://www.mysmartlife.eu/fileadmin/user_upload/deliverables/d4.1_baseline_report_of_helsinki_demonstration_area.pdf acta polytechnica ctu proceedings 38:183–189, 2022 1 introduction 2 background and current practices in smart cities 3 mysmartlife methodology: translating kpis into monitoring 3.1 kpi-driven assessment framework 3.2 mapping kpis into monitoring requirements 4 examples of application in the cities 4.1 nantes case 4.2 helsinki case 5 conclusions and lessons learnt acknowledgements references https://doi.org/10.14311/app.2022.33.0370 acta polytechnica ctu proceedings 33:370–376, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague experimental investigation on the mechanical behaviour of aac blocks for sustainable concrete masonry elena michelini∗, daniele ferretti, alice sirico, roberto cerioni university of parma, department of engineering and architecture, parco area delle scienze, 181/a, 43124 parma, italy ∗ corresponding author: elena.michelini@unipr.it abstract. to satisfy the increasing demand of energy efficient buildings, aac manufacturers are nowadays encouraged to produce blocks with ever lower densities. however, a compromise between energy-saving requirements and mechanical performances is needed to ensure structural safety, as well as an adequate structural durability. this paper reports a comprehensive experimental study on aac mechanical properties (compressive and tensile strengths, as well as fracture energy), and on their dependency from material density and moisture content. the collected data are compared with some well-known analytical relations taken from the literature, which are often used for the calibration of mechanical parameters required for mathematical and/or finite element modelling of aac load-bearing masonry, as well as of aac masonry-infilled framed structures. these comparisons highlight some critical issues in the formulation of analytical relations having a general applicability; however, it was found that rilem suggestions are appropriate for the considered aac productions, at least for densities greater than 400 kg/m3. keywords: autoclaved aerated concrete, mechanical properties, sustainable concrete masonry. 1. introduction the increasing interest in energy efficient buildings has promoted a growing use of innovative masonry products in the construction market. this tendency to search for new and non-standard masonry techniques is related to the widespread awareness that a large proportion of thermal losses in buildings is due to masonry walls. therefore, one of the most efficient strategies for the reduction of greenhouse gases produced by space heating is the adoption of products with increased thermal insulation performances [1–3]. among them, autoclaved aerated concrete (aac) is becoming a quite common solution for the realization of sustainable concrete masonry units (cmus), to be used in the construction of new buildings, as well as in the retrofitting of existing ones [4–10]. aac is a porous lightweight concrete whose cellular structure is generally obtained through a gas-producing chemical reaction of sand, lime, gypsum, and cement slurry, with the addition of an expanding agent (usually aluminium powder). other waste materials, such as fly ash, bottom ash, air-cooled slag, slate waste, glass or perlite waste, agriculture and industrial waste, etc. can be added to increase the sustainability of the production process [11–14]. thanks to the large amount of entrapped air ranging from 60% to 85% by volume the volume of the final product is up to five times that of the raw materials used in the production process. adequate strength and dimensional stability are obtained through autoclaving at high temperature and pressure, with the formation of a crystalline binder, called tobermorite [15]. thanks to the presence of small air voids (in the range of 0.1 − 1 mm) uniformly distributed in the cement paste, aac masonry blocks have many advantages in comparison with conventional concrete: lighter weight (typically from one-sixth to one-third of conventional concrete), lower transportation and building costs, excellent workability and easiness of laying, reduced construction times, remarkable acoustic and thermal insulation. besides its low thermal conductivity (which can reach 0.08 w/(mk), or even less for lower density values, see [16]), aac is characterized by an outstanding air-tightness, which further enhances building thermal performances, creating a comfortable living environment. furthermore, the use of aac, with its simple construction details, reduces energy losses due to thermal bridges at junctions [6]. the flip side of the coin is that the reduction of aac density, which improves thermal properties and reduces structural masses (and consequently the selfweight and the seismic forces acting on the building), exerts an unfavourable effect on material mechanical strengths. this could represent a limitation when aac blocks are used for the realization of load-bearing masonry walls, since in this case it is mandatory to find a compromise between satisfactory insulation requirements and minimum compressive strength values, according to standard codes [17]. for example, regardless of the type of adopted ma370 https://doi.org/10.14311/app.2022.33.0370 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 mechanical behaviour of aac blocks figure 1. sampling scheme for compression and flexural tests. sonry units, en 1998-1 [18] requires a minimum normalised compressive strength of 5 mpa, so to ensure adequate stability and robustness of masonry walls in seismic areas. even if not explicitly required by design standards, minimum tensile strength values (and consequently minimum compressive strengths, since these two properties are strictly related to each other) should be also provided for cladding and infill panels in framed structures, to limit crack formation under static and seismic loads. it is indeed well known that one of the most common drawbacks in using aac for the realization of cladding and infills is related to its limited fracture toughness and to its brittleness, which can reduce in some cases the durability of masonry panels, causing the appearance of premature and excessive damages [19, 20]. aim of this work is to provide a further insight on the relation between density and mechanical properties of aac. despite the large number of research works on this topic, a well-established relation between compressive strength and density has not yet been found, since available experimental data are referred to different raw materials in the admixture and to different autoclaving conditions (e.g., [3, 21–24]). on this point, another critical aspect is that temperature, pressure and curing time values are often not specified, despite their marked influence on the formation process of a stable form of tobermorite. a further complication relies on the fact that compressive strength is also dependent from the moisture content at time of testing, as well as from the possible treatments undergone by the specimen to smooth its surfaces (i.e. use of sandpaper, or of a water grinding machine). to better clarify these aspects, an experimental campaign is presented on aac specimens characterized by four different densities and four different moisture contents. the main mechanical properties, i.e. compressive and flexural tensile strengths, and the fracture energy, were determined according to relevant standards [25–27]. the best-known relationships between these properties and the material density, as taken from the literature [23, 24, 28], were then checked against the collected data. 2. experimental program 2.1. preparation of aac specimens all the specimens were cut from commercially available aac masonry blocks with nominal dimensions equal to 600 × 250 × 300 mm. the blocks were manufactured at the production plant starting from the following raw materials: sand with high silica content, cement, lime, water, gypsum, and aluminium powder as expanding agent. the blocks were autoclaved for 11 hours, at a temperature of 180 ◦ c and a pressure of 12 bars. four different types of commercial blocks were investigated, with nominal density equal to 300, 350, 480 and 580 kg/m3 (respectively indicated with symbols a, en, e, and s in table 1). the sampling scheme followed for compression and flexural tests is depicted in figure 1. as can be seen, 6 cubes with 100 mm side and three 50 × 50 × 200 mm prisms were cut from the bottom, middle and top part of each block (with respect to the direction of rising), by using a diamond blade cutter and a hacksaw for aac. three of these six cubes were used 371 e. michelini, d. ferretti, a. sirico, r. cerioni acta polytechnica ctu proceedings figure 2. test setup for the determination of: (a) compressive strength; (b) flexural tensile strength; (c) fracture energy. for the determination of the oven-dry density (db, dm and dt), while the remaining three were tested in compression (cb, cm, ct), according to [25, 29]. prismatic samples were instead tested in flexure (fb, fm, ft) according to [26]. for each examined density, the sampling was repeated on six blocks (so having a total of 144 cubes and 72 prisms). according to the adopted standards, before the tests, specimens were cured at laboratory conditions until the attainment of a moisture content equal to (6 ± 2) %. in order to investigate the effect of moisture content on mechanical strengths, additional samples (108 cubes and 54 prisms) were realized from 18 aac blocks with nominal density of 350 kg/m3. these specimens were subjected to different curing conditions so to achieve a moisture content respectively equal to (0 ± 2) %, (15 ± 2) % and an "upper limit" equal to the moisture content of the samples at their delivery to the laboratory (approximately ranging from 20 to 30%). before conducting the tests, flatness and parallelism requirements were checked; in some cases, it was necessary to smooth sample surfaces with sandpaper. regarding fracture energy determination, the only available standard for aac is represented by rilem recommendations [28], which suggests to carry out wedge-splitting tests. these tests are quite complex to be performed and require a completely different setup with respect to flexural tests. on the contrary, for standard concrete, fracture energy is usually determined by means of three-point bending tests on notched specimens. previous works (i.e. [22]) proved that this test method can be extended also to aac, providing results in accordance with those of wedgesplitting tests. for this reason, in this work fracture energy tests were carried out on notched prismatic samples having the same geometry as those tested in flexure, following the japanese code [27]. it was decided to extract only one prismatic specimen from the middle of each block, together with two cubes for the determination of the corresponding oven-dry density and compressive strength. three blocks were tested for each considered density and moisture content (also in this case, the latter was varied only for those blocks with a nominal density of 350 kg/m3). 2.2. determination of oven-dry density as already stated, three twin cubes extracted from each block (db, dm and dt according to figure 1) were used for the determination of the oven-dry density. to this aim, the specimens were dried in a ventilated oven at the temperature of (105 ± 5) ◦ c until a constant mass was reached. the oven-dry bulk density was then calculated as the oven-dried mass divided by the volume. the moisture content was determined as the ratio between the loss of mass during drying and the corresponding oven-dry mass. 2.3. compression tests according to [25], compression strength was calculated as the average of the results obtained on three standard cubes with an edge length of 100 mm, cut from the top, middle and bottom part of each block. compression tests were performed by using an instron 5882 press working under loading control (figure 2a), with a loading rate of 0.05 mpa/s, as suggested for masonry elements with an expected compressive strength lower than 10 mpa. loading was applied perpendicular to the direction of rise. 2.4. flexural tests aac modulus of rupture (tensile strength in bending, so-called mor) was obtained from prismatic samples subjected to three-point bending over a net span of 150 mm. an instron 8862 press was used to the scope, with a loading rate of 10 n/s (figure 2b). flexural tensile strength was determined as [26]: fct = 1.5 f l bf r h 2 f r (1) where f is the failure load, l is the net span, bf r and hf r are the specimen dimensions in correspondence of the cracked section. 372 vol. 33/2022 mechanical behaviour of aac blocks figure 3. evaluation of the average fracture energy for specimens with nominal density of 350 kg/m3 and moisture content ws of (0 ± 2) % by considering: a) the load p -deflection δ curve; b) the load p-cmod curve, according to [27]. density class thermal conductivity nominal density oven-dry density compressive strength flexural tensile strength (w/(mk)) (kg/m3) (kg/m3) (mpa) (mpa) a 0.07 300 ± 50 295 (0.007) 1.93 (0.04) 0.65 (0.13) en 0.08 350 ± 50 346 (0.008) 2.56 (0.03) 0.77 (0.08) e 0.11 480 ± 50 506 (0.002) 3.94 (0.08) 1.13 (0.05) s 0.13 580 ± 50 588 (0.005) 5.49 (0.08) 1.44 (0.05) table 1. properties of aac specimens for different density classes, with (6 ± 2) % moisture content. 2.5. fracture energy tests the test setup for the determination of fracture energy was similar to that adopted for flexural tests (figure 2c). however, in this case, the specimens were notched at their mid-length (with a notch depth equal to 0.3 times the beam depth) and the tests were performed under crack mouth opening displacement (cmod) control, with a loading rate of 0.01 mm/min. the specimens were also instrumented with a linear variable displacement transducer (lvdt) for the measurement of midspan deflection. fracture energy gf was calculated as [27]: gf = 0.75 w0 + w1 alig (2) w0 being the area below load-cmod curve up to specimen failure, w1 the work done by the specimen deadweight and by that of the loading jig, and alig the area of the broken ligament. the applicability of this relation (which was originally developed for standard concrete) to aac was preliminary checked by comparing the so obtained fracture energy with that calculated as the total work of fracture given by the area under the complete load-midspan deflection curve, divided by the ligament area [30]. the deadweight of the specimen and of the loading arrangement was taken into account also in this case. as an example, fracture energy deduced from the load-displacement curve (where the midspan deflection was that measured by the lvdt) and that calculated from the load-cmod curve according to the japanese standard [27] is reported in figure 3 for three prismatic samples with nominal density of 350 kg/m3 (class en), and a moisture content of (0 ± 2) %. 3. results and discussion 3.1. oven-dry density for the four density classes considered in this study, the average oven-dry density was found to be equal to 295, 346, 506 and 588 kg/m3, with a coefficient of variation (cov) reported in brackets in the table below 0.01. these densities were almost coincident with the nominal values provided by the manufacturer, as can be seen in table 1. the declared nominal values of thermal conductivity are also reported in the table for each density class. 4. compressive strength the average cube compressive strength is summarized in table 1 for all the considered density classes, for a predefined moisture content equal to (6 ± 2) %. as can be seen from figure 4a, compressive strength increased with density due to the corresponding reduction in material porosity, but the relation between the two variables was found to be non-linear. the experimental relation between compressive strength and density was checked against some wellknown analytical expressions available in the literature. as can be seen, experimental data laid within the range suggested by rilem recommended practice [28], except for the lower density values. in this case, experimental compressive strengths were higher 373 e. michelini, d. ferretti, a. sirico, r. cerioni acta polytechnica ctu proceedings figure 4. comparison between experimental data and analytical relations available in the literature: a) cube compressive strength fcc vs. density ρ; b) dimensionless cube compressive strength fcc/fcc (6 %) vs. moisture content ws. figure 5. (a) comparison between experimental data and analytical relations available in the literature in terms of flexural tensile strength fct vs. cube compressive strength fcc; (b) experimental relation between fracture energy gf and density ρ, for a moisture content of (6 ± 2) %.. than the predicted ones. in the same graph, the relations proposed by argudo [24] and by chen et al. [23] were also plotted. in these relations, which were obtained from the best-fitting of experimental data collected by the authors, the compressive strength is expressed as a function of the bulk density corresponding to a moisture content of 10% (ρ10 %), by assuming that the ratio ρ10 %/ρdry is equal to 1.1. the comparison between the experimental data collected in this study and the two considered analytical relations shows that a clear relation between compressive strength and density has not yet been found for aac. figure 4b highlights the influence of moisture content at time of testing on the dimensionless compressive strength. it can be seen that the maximum strength was found at the end of the drying process. even if standard codes allow to perform compressive tests on dried specimens (by introducing a correction factor on the so obtained strength), the actual moisture content normally reached by aac blocks in one or two years in external constructions is about 4−6 % [30]; for this reason, it seems more reasonable to provide the compressive strength value corresponding to that condition. in addition, the compressive strength of specimens tested in dry conditions may be influenced by the required thermal treatment. as can be seen in figure 4b, experimental normalized strength at dried condition was found to be significantly lower than that suggested by rilem recommended practice [28], while experimental values for other moisture contents fell inside the proposed range. 4.1. flexural tensile strength and fracture energy experimental values of flexural tensile strength coming from three-point bending tests are summarized in table 1. as for ordinary concrete, flexural tensile strength is generally expressed as a function of concrete compressive strength, instead of density. as can be seen in figure 5a, the relation suggested by 374 vol. 33/2022 mechanical behaviour of aac blocks rilem [28] was found to well fit the experimental data collected in the performed experimental campaign, while the semi-empirical equations suggested more recently by argudo [24] and by chen et al. [23] had a significantly lower precision, especially for higher compressive strengths. this is a further evidence of the limits of semi-empirical relations available in the literature, which suffer the dependency of aac mechanical properties from the raw materials adopted in the production process, the autoclaving conditions and the moisture content of the specimens at time of testing, and consequently they could hardly be generalized. a possible strategy to overcome these limitations could be the formulation of different relationships associated to distinct production processes, and to fix a predefined value of moisture to be reached before testing. even if tensile strength is the main parameter governing crack formation, it can be useful also to know the fracture energy, since it affects crack propagation within the material. when masonry behaviour is analysed through finite element analyses, the knowledge of fracture energy is necessary for the calibration of a proper block cohesive law. however, the dependency of fracture energy with material density has been little studied so far. a possible relation based on the data collected in this experimental program is shown in figure 5b. this relation well fit also the average of the data collected by ferretti et al. [22] in a previous experimental work, on a different aac production with a density of about 550 kg/m3. however, a further validation with other experimental programs is required. 5. conclusions to reduce thermal losses in buildings, an increasing attention is being paid to the reduction of cmus thermal conductivity and, consequently, of their density. however, minimum mechanical properties should be guaranteed also for lower density blocks, so to ensure adequate structural performances and the compliance with durability requirements (i.e. limiting crack formation in infills and claddings). this work focuses on the mechanical characterization of aac blocks belonging to four density classes. based on the obtained experimental results and their comparison with some semi-empirical relations available in the literature (calibrated on other aac productions), the following conclusions can be drawn: • an almost linear relation can be inferred between thermal conductivity and density, as well as between fracture properties (fracture energy) and density, while the relation between compressive strength and density seems to have an exponential form; • the relation between mor and compressive strength is almost linear; • significant variations may be obtained in compressive strength for a given density value, mainly due to different production processes and material curing; • eco-mechanical indexes, which are often used as synthetic indicators of environmental and mechanical performances for concrete (see, e.g. [4]), are hardly applicable in the considered case, unless modifying the definition of the reference environmental data (which are typically related to the production process, i.e., global warming potential and embodied energy, and are consequently quite similar for the four density classes). on the contrary, the major advantages of low density blocks are related to energy savings during the use phase of the building life cycle, which should be somehow taken into account. acknowledgements this work is part of a research founded by ekoru s.r.l., within pon national project n. f/090017/00/x36, cup b18i17000450008. authors gratefully thank dr. f. talento, arch. a. riva and dr. l. bergamonti for their valuable help. references [1] uygunoğlu t and kec̨ebas̨. lcc analysis for energy-saving in residential buildings with different types of construction masonry blocks. energy and buildings 43(9):2077-2085, 2011.https: //doi.org/10.1016/j.enbuild.2011.04.011. 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cc-by 4.0 licence published by the czech technical university in prague a guideline to support the use of off-site solutions for façade retrofitting through bim-enabled processes marco cucuzza∗, andrea giuseppe di stefano, giuliana iannaccone, gabriele masera politecnico di milano, department of architecture, built environment and construction engineering, via ponzio 31 – 20133 milano, italy ∗ corresponding author: marco.cucuzza@polimi.it abstract. the architecture, engineering and construction sector requires an intense innovation process to reduce costs, intervention times, and improve energy performance. this is particularly true to increase the rate of deep renovations of the existing european building stock and therefore support the eu’s 2050 decarbonisation targets. in this scenario, prefabrication can be considered a game-changer for the construction industry: on the one hand, it enables the adoption of an industrial mindset to the design, manufacturing, and installation of façade components, with the related advantages in terms of time, cost and quality; on the other, it allows for the customisation of components thanks to digital, bim-based design and fabrication tools. however, its adoption in retrofit operations is still limited due, among other factors, to a lack of practical experience and a limited number of actual demonstration cases. this paper introduces an operative guideline to support a more widespread use of prefabricated thermal insulation components for the energy retrofit of existing buildings in the framework of a bimenabled design and construction process. the guideline highlights the information flow and the role of each actor at every stage of the design and delivery process. the proposed guideline is finally tested through its application to a case study to show the feasibility of the process and the advantages deriving from the adoption of an industrialised approach to façade retrofit in terms of faster installation times. keywords: energy retrofit, mmc, dfma, off-site, bim. 1. introduction as widely discussed in literature, the construction sector needs an intense innovation process to reduce costs, intervention times, and energy needs [1]. achieving a significant energy renovation of the existing building stock would lead to an 80 % reduction in energy demand in 2050 compared to 2005 levels [2]. in this scenario, the european union faces a double challenge: to at least double the annual energy renovation rate by 2030, and to foster deep energy renovations (i.e., renovations that reduce energy consumption by at least 60 %) [3]. modern methods of construction (mmc) are widely regarded as a promising solution to these challenges and, more generally, to many shortcomings of the construction industry. mmc are a broad term that encompasses different dimensions of innovation, such as off-site construction and the related digital tools and techniques [4]. off-site construction is an approach to construction projects that seeks to move the construction process away from the site and take advantage of manufacturing approaches and standardisation efficiencies. the advantages of the mmc adoption are mainly the increase of quality (+ 30 %) and safety of workers (+ 80 %) together with a reduction of costs (20 %), time (20 to 50 %) and energy consumption (30 %) due to better management of resources in offsite activities instead of the on-site ones (70 %) [4–6]. the uncertainty of a dynamic working environment, the highly fragmented supply chain of smes, the need for skilled workers, the limited digitalisation, the shortterm thinking, and a political/governmental/cultural opposition are the main barriers to the mmc adoption. in contrast, horizontal integration of networks, endto-end digital integration, vertical integration, and networked manufacturing systems are the three keys to implement industry 4.0 in buildings [7]. the energy retrofit of existing buildings requires a tailored approach for each situation, because of the variety, among others, of construction periods, building technologies, materials, local climatic conditions, and site constraints. additional differentiating issues are local regulations and budget concerns. in parallel to this complex task, construction companies have a limited propensity to innovation, resulting in a productivity gap. in fact, their production rate grows only 1 % per year, compared to the medium growth rate of 2.8 % of other sectors. according to literature [2], productivity can be boosted by 50–60 % by implementing regulation, contract, design, supply chain management, and improving on-site execution 323 https://doi.org/10.14311/app.2022.38.0323 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en m. cucuzza, a. g. di stefano, g. iannaccone, g. masera acta polytechnica ctu proceedings by technology innovations and reskilling workers. these considerations show the market opportunities laying in the development of off-site prefabricated components for envelope retrofitting, reducing building energy demand and simplifying installation procedures while minimising the discomforts for the occupants. from this point of view, many research projects [8– 12] focused on developing new retrofit systems based on totally off-site or hybrid on-site/off-site technologies, while others, such as bim4eeb [13], worked on the related bim-based information structure. a core aspect for the large-scale implementation of mmc is their integration in a bim approach and the consequent adoption of rich digital communication that can increase interoperability levels, quality of information, and collaboration between stakeholders [14]. off-site automated activities, facilitated by a bim approach, have the potential to deliver the most significant growth in productivity in the construction sector [6]. for these reasons, various guidelines, design and assessment tools were developed specifically for new construction [15–17], but their application to existing buildings has been less explored so far. some research projects aim to streamline the decisionmaking process and integrate bim-based toolkits to facilitate the early-stage selection of retrofit technologies [18, 19]. this paper presents a general-purpose guideline on the bim implementation for prefabricated thermal insulation components, developed in the framework of the european research project bim4eeb. the guideline aims to go beyond the specific aspects of each construction technology, while focussing instead on the information exchange across the design, fabrication and delivery process of the off-site panels and between stakeholders (among others, architect, engineer, façade designer, manufacturer, supplier and general contractor). 2. methodology to develop the guidelines, the first step was an analytical review of previous research projects and off-site products already available on the market. the design, production, installation process, and the information exchange were investigated through literature, reports, available technical information and interviews. in the second step, the information gathered was used to produce a scheme of the steps and information exchange taking place across a general, technologyneutral, off-site process. the information flow defines the path of data through each stage, the actors involved and the necessary tools. in the third step, an analytical guideline was developed, in a form that would be useful both to designers interested in using prefabricated façade panels and to companies developing such solutions. the guideline is presented in the form of a process matrix, highlighting the information flow at each stage of the process and providing information about the actions required from the relevant actors. finally, the guideline has been tested on a real-life case study building model to assess the feasibility of the information flow and the applicability of the different process stages. 3. guidelines the guidelines support designers in identifying the information required at each stage of the process, streamlining the information exchange and enabling more effective communication with the other stakeholders involved (e.g., manufacturer and client). on the other hand, a guideline is also useful for manufacturers who plan to enter, or already are in, the field of off-site panels for the retrofit market, which is often characterised by an ad-hoc approach to each application because of its relatively short history. in this case, manufacturers can organise their technical information in a way that is compliant with the guideline and can also activate early engagement strategies to support the adoption of their products in retrofit operations. the information gathered from existing prefabrication technologies, eu-funded projects [8–10, 20] and interviews with european manufacturers made it possible to arrange the whole design-to-construction process of façade panels according to phases with general applicability, independent from the specific construction technology. the following diagram (figure 1) presents each phase with its purpose (task) and its outcome, i.e., the information it produces for the following steps of the process. this articulation will constitute the basis of the proposed framework. the framework (table 1, appendix a) shows the inputs, actions, and outputs in a clear, step-by-step process matrix, particularly suitable to highlight the correlations between the various steps and the related project phases. the steps column (second column) describes the general scope for each stage. then, the main objectives (third column) are identified, together with the required input activities (fourth column) carried out by other design team members; these activities are preparatory and necessary for those more strictly related to the panelisation activities. finally, under the output activities column (fifth column) and the data/model development column (sixth column), the required actions are summarised both from the process and the modelling point of view. the process matrix is discussed among european manufacturers and design teams to suggest a step ahead in the knowledge-based engineering [21] adoption compared to the existing professional procedure and the scientific literature [22, 23]. compared to other workflows, the proposed one started from the adoption of “panels place orders” during the concept design, allowing parametrical tools to generate various possible solutions based on the façade designer expertise, the manufacturer limitation in the production phase 324 vol. 38/2022 a guideline to support the use of off-site solutions . . . figure 1. information flow for panelisation process. and the site-specific constraints. the proposed solutions are meta-technological products – underpinned on real ones – avoiding the risk of redesigning the entire façade after the manufacturer engagement and, simultaneously, allowing the progressive information enrichment throughout the process thanks to the bimbased design process. stages 1 and 2 refer to the acquisition of information about the existing building, while the design phases (stages 3–5) concern the project of the panels increasing the level of details thanks to the iterations that the parametrical instruments enabled (design optioneering). finally, construction, building use and end of life (stages 7–9) describe the transmission of the data-rich model between the designer and the manufacturer, avoiding the potential loss of information with consequences in the management and disposal phases. at the same time, returning an informed model from the manufacturer to the designer and contractor makes it possible to integrate the assembly sequences smoothly into the site activities. this holistic perspective can significantly reduce construction time, allowing parallel activities on the actual construction site and in the factory. 3.1. case study the framework for the design of prefabricated thermal insulation components was tested on a case study to verify that the steps of the process and the information flow can be used in practice. a simplified version of one of the demonstration sites of the bim4eeb research project – a multi-storey multi-owner building located in monza (italy) – was used for this purpose. the exercise focuses on the first stages of the framework (from 1 to 5), demonstrating its effectiveness and how to select the configuration that best meets the requirements set by the client. the output (design of the prefabricated panels fitting the specific case study) will contain the necessary information to interface downstream with production and site operations (stages 6 to 9). • stages 1, 2: the data flow was defined and the as-is bim model was produced based on a survey. • stage 3: the first phase of the optioneering process was performed about the geometrical aspects of the façade panels. meta-technological objects were used as placeholders to reproduce different panelling hypotheses evaluated with the stakeholders. each time one panelling option was analysed, an automated optimisation (considering specific dimensional constraints) was conducted using a bespoke script, following the goal of maximising the number of equal panels, minimising the total number of panels, and reducing or eliminating unique panels. the software then provided a datasheet identifying individual elements (figure 2). • stage 4: starting from the first optioneering iteration results, several automated parametrical tools performed detailed analyses. once the dimensional and technological information were included, it was possible to evaluate construction details, such as corner solutions. these are crucial topics to be assessed in the early stage because they influence the whole retrofitting operations behaviour in terms of heat loss, thermal bridges, and aesthetics. each alternative was fully defined, and spreadsheets were exported for client evaluation and selection of the technology. • stage 5: the chosen panels were exported back into the bim environment, and a complete ifc model was shared with the manufacturer for validation (early-stage contractor and supply chain engagement). after an early-stage discussion with the manufacturer, the panel layers were designed in detail, defining the specific materials, thickness and other properties, anchors positioning, and dimensions. a detailed structural analysis was also performed in the bim environment, automatically defining the positions of anchors in the bim model through a dedicated script. once the model included the panels and the anchors, it was possible to proceed with the detailed energy analysis, eval325 m. cucuzza, a. g. di stefano, g. iannaccone, g. masera acta polytechnica ctu proceedings figure 2. as-is bim model and first panelisation phase with meta-technological panels. the total number of panels and the number of different panels considered is automatically displayed. figure 3. perspective view of the bim model with the selected panelisation. uate thermal bridges, and check compliance with local regulations. at the end of stage 5, a fully verified and integrated bim model was ready to be shared with all the stakeholders, particularly downstream with the manufacturer. in this whole process, and particularly in stage 5, the manufacturer’s ability to use and return an information-rich model is fundamental to ensure the smooth application of the guidelines. the capability to receipt, manage, transform and transmit the information required for each actor during each stage was verified thanks to the check of the final ifc model and the other additional documents (bill of panels, manufacturing drawings, installation drawings and technical documentation), although it was not possible to test the guideline with the actual production and installation of the panels. 4. conclusions the paper aims to propose a guideline to support the design of prefabricated thermal insulation components for facades of existing buildings within a bim-based approach. this guideline can encourage wider market uptake of off-site systems for retrofit operations, contributing to maximising efficiency in building renovation, reducing the renovation working time, enabling more accurate building quality control, and implementing a bim-based renovation business for construction companies as part of the european transition towards a digital built environment. the guideline presents, in a systematic way, the digital information flow necessary to support the design and delivery of prefabricated façade panels, with the specific data and involved actors correlated to the various stages of the process. thanks to the abstraction of current practices and procedures, the guideline has broad applicability and is not technology-dependent and country-specific. finally, it was tested on a case study to assess the practical feasibility of the process, including aspects such as the possibility to retrieve the necessary data and the information flow between different software and tools. the guideline represents a relevant tool to support designers who intend to adopt prefabricated façade systems in their retrofit projects, clarifying what information is required at each step of the process, from/to whom it should be provided, and what decision-making tools can be used. furthermore, the guideline also supports construction com326 vol. 38/2022 a guideline to support the use of off-site solutions . . . panies/manufacturers of façade systems, anticipating what information should be made available to the other actors of the process to ensure a smooth integration of the prefabricated panels in the bim-based workflow early-stage engagement of the industry when possible. the integration of rules and constraints provided by all the actors – including manufacturers, architects, construction site managers and façade designers – at the beginning of the process, together with the bim willingness to increment information across the stages, can be a step ahead in the adoption of knowledge-based engineering in the design for manufacture of prefabricated façades. acknowledgements this paper builds upon results from the bim4eeb project, which has received funding from the european union’s horizon 2020 research and innovation programme under grant agreement no 820660. it reflects only the author’s view and that the commission is not responsible for any use that may be made of the information it contains. references [1] r. hammond, n. o. nawari, b. walters. bim in sustainable design: strategies for retrofitting/renovation. in computing in civil and building engineering (2014), pp. 1969–1977. 2014. https://doi.org/10.1061/9780784413616.244 [2] mckinsey global institute. reinventing construction through a productivity revolution, 2017. 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base for future work. • accurate survey of the building (laser scanner or other methods); • collect technological, historical and occupancy information; • 3d model development and information collection; • develop as-is model of the building, including basic information (structural grid, thermal transmittance, etc.). • early assessment of opportunities for mmc adoption. • as-is bim model. 3 c on ce pt d es ig n • develop early-stage design coherently with the project scope and objectives. • explore feasibility of panelisation strategy; • define geometrical alternatives. • as-is bim model; • knowledge-based constraints. • geometrical and design exploration; • use meta-technological objects to generate design optioneering; • develop parametric placeholders panels to understand possible issues. • include information such as massing (panels placeholders), meta-technological information and basic information (structural grid, thermal transmittance). 329 m. cucuzza, a. g. di stefano, g. iannaccone, g. masera acta polytechnica ctu proceedings stage step(s) objective(s) input mmc-related panelisation activities data/model development 4 p re lim in ar y d es ig n • develop and test the prefabricated solution; • early supply chain involvement. • early evaluation of technical alternatives. • client feedback from stage 3 • acquire performance targets, in particular u-value. • technological alternatives exploration for reaching project target; • generate datasheets from objects for approval of functional, environmental and finishes requirements; • identify supply chain and producers for technology evaluation. • develop multiple models with data sheets about technical alternatives. 5 d ev el op ed d es ig n • design analysis and calculations. • confirm full conformity to scope and objectives of the project; • early engagement of contractors and supply chain. • client feedback from stage 4; • perform energy simulations of the existing building to understand possible issues for the retrofitting; • perform structural simulations of the existing building to understand possible issues for the retrofitting; • perform lca simulations of the existing building to understand possible issues for the retrofitting. • add more details to bim objects (both geometry and data); • validate panelling technological solutions through early contractor and supply chain engagement; • use analytical tools and calculations to compare technological alternatives; • generate detailed part and whole models for different disciplines for early coordination; • final choice of the panel technology; • parametric time and cost validation. • include details of the chosen technical solution in the bim model. 6 d et ai le d d es ig n • production design for fabrication; • digital prototype of the panels. • develop full model with information from the manufacturer. • integrate supply chain information about resources and time. • develop overall construction programme schedule and assembly sequencing; • incorporate inputs from supply chain into the bim model; • develop fabrication and installation sequences, resource management plan, etc.; • generate digital prototypes of the panels to verify the construction process; • detailing of panels (connections, special parts, etc.). • construction model including highly accurate information from supply chain. 330 vol. 38/2022 a guideline to support the use of off-site solutions . . . stage step(s) objective(s) input mmc-related panelisation activities data/model development 7 7. 1 p re -c on st ru ct io n • accurate fabrication model. • enable fabrication of components. • accurate bim model from stage 6. • generate shop drawings for fabrication from models or integrate fabrication information from the manufacturer into the bim models. 7. 2 o nsi te c on st ru ct io n • training; • virtual building. • enable construction in line with scope and objectives of the project; • inform users about construction activities (time and location). • construction model updated with pre-construction information. • track construction activities and resources based on planned programme and planned assembly sequence; • validate installation on-site and update as-built model. • interface and communicate with users about construction activities. • as-built model incorporating information from previous stages and construction activities. 8 b ui ld in g u se • life cycle costs; • facility management attributes; • maintenance program. • represent the as-built asset; • produce ordinary maintenance plan. • maintenance information from supply chain; • as-built model from stage 7. • integrate as-built model with the facility management system; • integrate information about maintenance in the as-built bim model. • as-built models are up-to-date and accessible to the property management and tenants; • represent the building in use, including dynamic feedback information from bms and fm (digital twin). 9 e nd of l ife • program end of life. • produce end of life actions plan. • digital twin model from stage 8. • develop end of life support documentation. • the model is up-to-date and will continue evolving. table 1. process matrix for energy retrofit with off-site façade panels. 331 acta polytechnica ctu proceedings 38:323–331, 2022 1 introduction 2 methodology 3 guidelines 3.1 case study 4 conclusions acknowledgements references a appendix – table 1 my title https://doi.org/10.14311/app.2022.36.0237 acta polytechnica ctu proceedings 36:237–243, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague structural reliability of hollow core slabs considering compressive membrane action thomas thienpont∗, ruben van coile, wouter de corte, robby caspeele ghent university, department of structural engineering and building materials, technologiepark-zwijnaarde 60, 9052 ghent, belgium ∗ corresponding author: thomas.thienpont@ugent.be abstract. compressive membrane action can considerably improve the load bearing capacity of concrete slabs and beams in case of excessive loaded due to an accidental event. currently, only limited research has been focusing on compressive membrane action in prestressed concrete elements, or on concrete elements with large cavities, such as precast concrete hollow core slabs. therefore, a novel real-scale test setup has been developed in order to assess this effect in precast hollow core slabs, and how it can enhance the load-carrying capacity in accidental events. in parallel with these tests, a numerical finite element model has been developed in order to perform a more detailed structural analysis of this phenomenon, and to study the influence of various input parameters. the details of this test setup are briefly explained, and some relevant experimental test results are provided. considering the experimental findings and validated numerical model, this contribution aims to quantify the influence of compressive membrane action on the structural reliability of precast concrete hollow core slabs. to this end, probabilistic models for the most important material and geometric variables are gathered, and the structural reliability is assessed using latin hypercube sampling. overall, the results indicate that considering the formation of compressive membrane action strongly influences the variability of the ultimate load-carrying capacity of precast concrete hollow core slabs. keywords: cost optimization, structural reliability, target reliability. 1. introduction when an axially restrained concrete slab is excessively loaded (e.g. due to an accidental event), compressive membrane action (cma) can be activated, which can considerably enhance the load-carrying capacity. this can delay or even prevent a progressive collapse, increase the structural reliability and consequently increase the robustness of concrete structures. the general principle of cma in an axially restrained slab is illustrated in figure 1. although this effect was already investigated by other researches [1–4], most investigations focus on reinforced concrete members. so far, very little attention has been paid to the role of membrane action in prestressed concrete members. furthermore, only limited research has been focusing on cma in concrete members with large cavities, such as hollow core (hc) slabs. hc slabs are nowadays very commonly used, with the annual global production of these elements reaching 50 million m2 [5], and a better understanding of how cma could affect the behaviour of these slabs is needed. in a recent study, the beneficial effect of cma on the load-carrying capacity of precast hc slabs has been demonstrated experimentally [6]. in parallel, a numerical finite element (fe) model was established in order to model the structural behaviour (i.e. compressive membrane action) as per the investigated test program. in this contribution, the details of the experimental test setup and the numerical model are briefly discussed, and some relevant experimental test results are provided. next, the numerical model is employed to perform a stochastic fe analysis of the ultimate load-carrying capacity. herewith, this paper undertakes some first steps to quantify and incorporate the uncertainties which are associated to the development of compressive membrane action in precast hc slabs exposed to excessive loading. 2. novel experimental test setup the beneficial effect of cma on the load-carrying capacity of hc slabs was demonstrated in an experimental study by [6], as illustrated in figure 2. herein, three tests were performed: two axially restrained tests (labelled hc1 and hc2) and one unrestrained test (labelled hc3). the tests were carried out on hc slabs casted by a local belgian producer, with a length of 10.4 m and a 1200 mm × 320 mm cross section (see figure 3). in both the restrained and the unrestrained tests, the slabs were loaded in a fourpoint bending setup. the slabs were supported by two roller supports at a distance of 1 m from the slab end (resulting in a free span of 8.4m) and the loads were applied at 1/4th of the span length. in the restrained tests, the horizontal outward displacements of the hc slab were restrained by four concrete blocks (1300 mm 237 https://doi.org/10.14311/app.2022.36.0237 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en t. thienpont, r. v. coile, w. d. corte, r. caspeele acta polytechnica ctu proceedings figure 1. basic principle of compressive membrane action. effect in real-scale precast hc slabs. in parallel, a numerical finite element (fe) model has been established in order to model the structural behaviour (i.e. compressive membrane action) as per the investigated test program. in this contribution, the details of the experimental test setup and the numerical model are briefly discussed, and some relevant experimental test results are provided. next, the numerical model is employed to perform a stochastic fe analysis of the ultimate load-carrying capacity. herewith, this paper undertakes some first steps to quantify and incorporate the uncertainties which are associated to the development of compressive membrane action in precast hc slabs exposed to excessive loading. 2 novel experimental test setup to investigate the beneficial effect of compressive membrane action on the load-carrying capacity of prestressed hc slabs, an experimental test set-up has been developed as illustrated in figure 2. three tests were performed: two axially restrained tests (labelled hc1 and hc2) and one unrestrained test (labelled hc3). the tests were carried out on hc slabs casted by a local belgian producer, with a length of 10.4 m and a 1200 mm x 320 mm cross section (see figure 3). in both the restrained and the unrestrained tests, the slabs were loaded in a four-point bending setup. the slabs were supported by two roller supports at a distance of 1 m from the slab end (resulting in a free span of 8.4m) and the loads were applied at 1/4th of the span length. in the restrained tests, the horizontal outward displacements of the hc slab were restrained by four identical concrete blocks (1300 mm × 400 mm × 1600 mm) that were anchored in the laboratory strong floor. additionally, uplifting of the slab ends was prevented by two steel hem160 profiles, which were anchored to the lab floor with two high-grade steel rods. this particular setup allowed for the formation of significant compressive membrane forces inside a real-size hc slab. in the unrestrained test, these end restraints were removed, and the outward expansion and rotation of the slab was not hindered. a) b) figure 2: restrained four-point bending test: (a) schematic overview (dimensions in mm); (b) picture of the actual test set-up figure 2. restrained four-point bending test: (a) schematic overview (dimensions in mm); (b) picture of the actual test set-up [6]. × 400 mm × 1600 mm) that were anchored in the laboratory strong floor. additionally, uplifting of the slab ends was prevented by two steel profiles, which were anchored to the lab floor with high-grade steel rods. this particular setup allowed for the formation of significant compressive membrane forces inside a real-size hc slab. in the unrestrained test, these end restraints were removed, and the outward expansion and rotation of the slab was not hindered. 2.1. test results during the tests, the vertical and horizontal displacements were measured using 9 displacement transducers. the applied forces and horizontal forces on the restraint blocks were also continuously measured using 6 load cells. in the two restrained tests, the slabs reached a capacity of respectively 298 kn and 269 kn per load application point, which was significantly higher than the unrestrained setup (capacity = 180 kn). the maximum horizontal reaction force measured in the restrained tests was 1221 kn and 779 kn for hc1 and hc2 respectively. in both tests, the horizontal restraint stiffness of approximately 250 kn/mm was observed. this values was derived from the measurements of the restraint forces and horizontal displacement of the slab ends. the load displacement graphs from the three tests are depicted in figure 4a. both restrained slabs collapsed after a brittle web shear failure, whereas the unrestrained slab exhibited a more ductile behaviour. even though the test setup was identical in test hc1 and hc2 a significant difference in the ultimate capacity and deflection at failure was observed. for further details on the experimental campaign the reader is referred to [6]. 238 vol. 36/2022 hollow core slabs structural reliability figure 3. hollow core slab section (dimensions in mm). test age (days) fc,cube (mpa) fc,cyl (mpa) fct (mpa) hc1 68 75.7 64.3 3.85 hc2 71 74.9 65.6 3.63 hc3 85 78.1 68.1 3.63 table 1. sample characteristics of material properties for concrete. strand/wire fp0.2 (mpa) fm (mpa) εu (%) ep (gpa) hc1 1789 1975 6.25 221.5 hc2 1782 1938 5.21 218.1 hc3 1627 1848 5.94 193.9 table 2. sample characteristics of material properties for prestressing steel. 2.2. material tests the tested hc slabs were made of c50/60 concrete. to obtain the actual mechanical properties of the concrete, compressive tests on cube samples (150 mm × 150 mm × 150 mm), cylinder samples (diameter 150 mm, height 300 mm) and flexural bending tests on prism samples (400 mm × 100 mm × 100 mm) were performed, see table 1. the prestressing strands (bottom reinforcement) have a specified characteristic tensile strength of 1860 mpa. the prestressing wires (top reinforcement) have a specified characteristic tensile strength of 1770 mpa. to obtain the actual mechanical properties, tensile tests were performed on the three types of prestressing reinforcement. from the obtained stress-strain curves, the values for the 0.2 % strain limit fp0.2, tensile strength fm, ultimate strain εu and young’s modulus ep were derived, see table 2. these values were used as inputs for the material models in the numerical models. 3. finite element modelling 3.1. model validation in parallel with the experimental tests, a detailed 3d numerical model was developed using the finite element (fe) software abaqus [7]. herein, nonlinear material models were employed to model the behaviour of the concrete and prestressing steel under large deformations in an adequate manner. the concrete was modelled using the abaqus concrete damaged plasticity model. herein, the stress-strain diagrams for concrete in compression and tension were based on en 1992-1-2 [8] and fib model code 2010 [9], respectively. the prestressing steel was modelled using a bilinear stress-strain diagram. for each model evaluation, the material properties in accordance with table 1 and 2 were used. in the numerical evaluations, 8-noded hexahedral (brick) elements (c3d8) and 2-noded linear truss elements (t3d2) are used to model the concrete and steel prestressing strands, respectively. assuming perfect bond, the prestressing strands are embedded within the concrete element, using the embedded constraint option in abaqus. the supports and end restraints were modelled to precisely correspond to the test setup. overall, the model is able to predict the experimentally obtained load-displacement curves, the failure mode and crack patterns with high accuracy, as shown in figure 4a and figure 4b, respectively. only for the case of hc1, the model somewhat underestimates the deflection and load at failure. 3.2. influence of concrete tensile strength as observed in the experimental tests, the ultimate capacity fq,max and horizontal restraint force fh,max of the hc slabs in the restrained setup exhibit a large variability. even though the test setup was identical for hc1 and hc2, the difference in the measured ultimate capacity was close to 30 kn. furthermore, the difference in the maximum horizontal restraint force was nearly 450 kn. these large differences could be attributed to the brittle shear tension failure mechanism, which was observed in both tests. as shown in figure 4b, this failure mode is characterized by a diagonal crack, which forms at mid-depth and propagates towards the support and the compression zone, 239 t. thienpont, r. v. coile, w. d. corte, r. caspeele acta polytechnica ctu proceedings 3 finite element modelling 3.1 model validation in parallel with the experimental tests, a detailed 3d numerical model was developed using the finite element (fe) software abaqus (dassault systemes, 2014). herein, nonlinear material models were employed to model the behaviour of the concrete and prestressing steel under large deformations in an adequate manner. the concrete was modelled using the abaqus concrete damaged plasticity model. herein, the stress-strain diagrams for concrete in compression and tension were based on en 1992-1-2 (cen, 2004) and fib model code 2010 (fib, 2012), respectively. the prestressing steel was modelled using a bilinear stress-strain diagram. for each model evaluation, the material properties in accordance with table 1 and 2 were used. in the numerical evaluations, 8-noded hexahedral (brick) elements (c3d8) and 2-noded linear truss elements (t3d2) are used to model the concrete and steel prestressing strands, respectively. assuming perfect bond, the prestressing strands are embedded within the concrete element, using the embedded constraint option in abaqus. the supports and end restraints were modelled to precisely correspond to the test setup. overall, the model is able to predict the experimentally obtained load-displacement curves, the failure mode and crack patterns with high accuracy, as shown in figure 4a and figure 4b, respectively. only for the case of hc1, the model somewhat underestimates the deflection and load at failure. (a) (b) figure 4: comparison of experimental results and numerical model: (a) load-displacement graphs; (b) crack pattern at failure in restrained test hc1 3.2 influence of concrete tensile strength as observed in the experimental tests, the ultimate capacity fq,max and horizontal restraint force fh,max of the hc slabs in the restrained setup exhibit a large variability. even though the test setup was identical for hc1 and hc2, the difference in the measured ultimate capacity was close to 30 kn. furthermore, the difference in the maximum horizontal restraint force was nearly 450 kn. these large differences could be attributed to the brittle shear tension failure mechanism, which was observed in both tests. as shown in figure 4.b, this failure mode is characterized by a diagonal crack, which forms at mid-depth and propagates towards the support and the compression zone, thus causing a brittle failure (araujo, loriggio, & da figure 4. comparison of experimental results and numerical model: (a) load-displacement graphs; (b) crack pattern at failure in restrained test hc1 [6]. thus causing a brittle failure [10]. in the absence of transverse reinforcement in thin-walled webs of the hc slab, the shear capacity of slabs depends only on the tensile strength of the concrete [11]. hence the concrete tensile strength might considerably influence the capacity of these slabs. to test this hypothesis, the numerical model for hc1 was re-evaluated for a range of input value for the concrete tensile strength between 3.5 mpa and 5 mpa, as shown in figure 5. from the obtained load-displacement curves, it is clear that the tensile strength of concrete has a very large influence on the cma capacity of the hc in the studied test-setup, especially, in the range between 3.5 mpa and 4.5 mpa. the above result highlights the importance of considering the stochastic nature of the concrete tensile strength in a probabilistic assessment of concrete slabs prone to shear failure. 4. probabilistic model in the following section, a stochastic evaluation of the previously presented fe model is performed, in order to obtain the distributions of the ultimate capacity fq,max of prestressed hc slabs in restrained and unrestrained conditions. in the restrained configuration also the distribution of the maximum membrane force fr,max (i.e. the horizontal force in the slab) is evaluated. given the significant computational power required for a single fe model evaluation, it would be impractical to perform such an analysis by evaluating a set of traditional monte carlo simulations. hence, a latin hypercube sampling scheme was used to keep the number of model evaluations to a minimum. furthermore, four key variables were selected, which are considered as the most influencing with respect to the determination of the ultimate load capacity, especially in case of the occurrence of cma. the key figure 5. influence of the concrete tensile strength fct on the load-displacement behaviour of restrained hollow core slabs. variables are the concrete compressive strength fc, the concrete tensile strength fct, the initial prestress σpi and the spring constant k corresponding to the horizontal restraints. the probabilistic models for concrete material properties and the initial prestress of the steel reinforcement are based on [12, 13], as presented in table 3. herein, the mean values for the concrete compressive and tensile strength are based the values of hc1, in accordance with table 1. as a first approximation, the concrete compressive strength and tensile strength have been considered independent. to study the influence of the axial restraint stiffness, two values for the spring constant k are considered, i.e. 50 kn/mm and 250 kn/mm. the latter value corresponds to 240 vol. 36/2022 hollow core slabs structural reliability property distr. µ cov concrete compressive strength fc ln 64.3 mpa 0.06 concrete tensile strength fct ln 3.85 mpa 0.3 initial prestress σpi n 1116 mpa 0.06 spring constant k of horizontal restraints ln case dependent (kn/mm) 0.25 table 3. probabilistic models for most important basic variables involved in the uncertainty quantification of the compressive membrane action in the investigated hollow core slabs. figure 6. lhs samples of restrained and unrestrained hollow core slabs; (a) the load-displacement curves (b) restraint force versus displacement curves. figure 7. (a) pdf of the ultimate capacity fq,max; (b) pdf of the ultimate restraint force fr,max. the stiffness of the restraints in the experiments. the value of 50 kn/mm was chosen to study the influence of a structure with a lower restraint stiffness on the development of cma in precast hollow core slabs. the probabilistic model for the spring stiffness is based on [14]. 5. results in total, three stochastic fe analyses were performed: two restrained cases and one unrestrained case. for each case, a set of 12 latin hypercube samples of the input variables were generated according to the method described in [15]. figure 6a depicts the numerically obtained load-displacement curves from the fe evaluations, for the unrestrained case and two restraint stiffnesses. the curves are displayed up to the point of failure (i.e. bending or shear tension failure). from the graph, it is clear that a higher restraint stiffness corresponds to a higher ultimate capacity fq,max and a reduced deflection at failure. the curves in figure 6b show the horizontal restraint force fr as a function of the midspan deflec241 t. thienpont, r. v. coile, w. d. corte, r. caspeele acta polytechnica ctu proceedings calculation distribution µ cov restrained 250 kn/mm ln 264.1 kn 0.084 restrained 50 kn/mm ln 227.1 kn 0.062 unrestrained ln 169.8 kn 0.048 table 4. probabilistic models for the ultimate load-carrying capacity fq,max. calculation distribution µ cov restrained 250 kn/mm ln 986.7 kn 0.279 restrained 50 kn/mm ln 352.5 kn 0.286 table 5. probabilistic models for the maximum horizontal restraint force fr,max. tion. in these curves, the restraint force increases monotonically, up to the point at which the ultimate capacity fq,max is reached. thereafter the restraint forces reduce rapidly (not shown in the graph). the graph illustrates that a higher restraint stiffness corresponds to a higher maximum restraint force fr,max. figure 7a shows the distribution of the ultimate load-carrying capacity fq,max, for each of the three considered cases, assuming a lognormal distribution. from this graph, it is clear that the variability of fq,max increases significantly with an increased restraint stiffness. additionally, figure 7b depicts the distribution of the maximum restraint force fr,max. this graph clearly illustrates the presence of a very large variability in the ultimate restraint force. considering the lhs samples in combination with the fe analyses, probabilistic models for the loadcarrying capacity and maximum restraint force of both the restrained and unrestrained cases are presented in table 4 and table 5 respectively. a lognormal distribution is assumed for both variables. the results in table 4 show that the coefficient of variation (cov) of the ultimate capacity fq,max is higher for the restrained cases, and increases with increasing axial restraint stiffness. table 5 on the other hand, indicates no significant difference in the value of the cov of the maximum horizontal restraint force fr,max for both restraint cases. 6. conclusions in the present paper, a stochastic finite element analysis was performed to study the influence of stiff lateral restraints on the load-displacement behaviour of precast concrete hollow core slabs. the finite element model was validated against a set of real-scale experimental tests, and was applied in combination with probabilistic models for the most influencing variables to execute latin hypercube simulations of the loadcarrying capacity. based on the random simulations, probabilistic models for the load carrying capacity and horizontal restraint force are established. in general, it can be concluded that the presence of lateral restraints strongly influences the variability of the ultimate load-carrying capacity. the coefficient of variation (cov) of the ultimate capacity increases from 0.048 in the unrestrained case to 0.084 for the highly restrained case with an axial restraint stiffness of 250 kn/mm. furthermore, a high cov of the horizontal restraint force was observed. this value was observed not to change significantly with respect to the two stiffnesses of the surrounding structure considered. acknowledgements this research has been made possible through funding from the research foundation flanders (fwo), which is gratefully acknowledged (grant number 11f5520n). references [1] l. k. guice, e. j. rhomberg. membrane action in partially restrained slabs. aci structural journal 85(4), 1988. https://doi.org/10.14359/2517. [2] g. i. b. rankin, a. e. long. arching action strength enhancement in later ally-restrained slab strips. proceedings of the institution of civil engineers structures and buildings 122(4):461-7, 1997. https://doi.org/10.1680/istbu.1997.29834. [3] k. thoma, f. malisia. compressive membrane action in rc one-way slabs. engineering structures 171:395-404, 2018. https: //doi.org/10.1016/j.engstruct.2018.05.051. [4] f. j. vecchio, k. tang. membrane action in reinforced concrete slabs. canadian journal of civil engineering, 17:686-697, 1990. https://doi.org/10.1139/l90-082. [5] k. s. elliott, c. k. jolly. multi-storey precast concrete framed structures. john wiley & sons, oxford uk, 2013. [6] t. thienpont, w. de corte, r. van coile, et al. compressive membrane action in axially restrained hollow core slabs: experimental investigation. structural concrete, 2022. https://doi.org/10.1002/suco.202100684. [7] dassault systemes. abaqus 6.14: abaqus/cae user’s guide. providence, usa, 2014. [8] cen. en 1992-1-2. eurocode 2 design of concrete structures. part 1-2: general rules structural fire design. brussels. https://www.phd.eng.br/wp-conte nt/uploads/2015/12/en.1992.1.1.2004.pdf. 242 https://doi.org/10.14359/2517 https://doi.org/10.1680/istbu.1997.29834 https://doi.org/10.1016/j.engstruct.2018.05.051 https://doi.org/10.1016/j.engstruct.2018.05.051 https://doi.org/10.1139/l90-082 https://doi.org/10.1002/suco.202100684 https://www.phd.eng.br/wp-content/uploads/2015/12/en.1992.1.1.2004.pdf https://www.phd.eng.br/wp-content/uploads/2015/12/en.1992.1.1.2004.pdf vol. 36/2022 hollow core slabs structural reliability [9] fib. fib model code 2010, final draft volume 1. comité euro-international du béton, 2010. http://repositorio.unan.edu.ni/2986/1/5624.pdf. [10] c. a. m. araujo, d. d. loriggio, j. m. m. n. da camara. anchorage failure and shear design of hollow-core slabs. structural concrete 12(2):109-19, 2011. https://doi.org/10.1002/suco.201000024. [11] w. derkowski, m. surma. prestressed hollow core slabs for topped slim floors theory and research of the shear capacity. engineering structures 241, 2021. https: //doi.org/10.1016/j.engstruct.2021.112464. [12] jcss. probabilistic model code part 3.1: concrete properties. in the joint committee on structural safety, 2000. [13] jcss. probabilistic model code part 3.4: static properties of prestressing steel. in the joint committee on structural safety, 2000. [14] r. caspeele, d. gouverneur, r. van coile, et al. structural reliability of concrete slabs considering tensile membrane action. in safety, reliability and risk analysis, p. 2713-20, 2013. https://doi.org/10.1201/b15938-410. [15] a. olsson, g. sandberg, o. dahlblom. on latin hypercube sampling for structural reliability analysis. structural safety 25(1):47-68, 2003. https://doi.org/10.1016/s0167-4730(02)00039-5. 243 http://repositorio.unan.edu.ni/2986/1/5624.pdf https://doi.org/10.1002/suco.201000024 https://doi.org/10.1016/j.engstruct.2021.112464 https://doi.org/10.1016/j.engstruct.2021.112464 https://doi.org/10.1201/b15938-410 https://doi.org/10.1016/s0167-4730(02)00039-5 acta polytechnica ctu proceedings 36:237–243, 2022 1 introduction 2 novel experimental test setup 2.1 test results 2.2 material tests 3 finite element modelling 3.1 model validation 3.2 influence of concrete tensile strength 4 probabilistic model 5 results 6 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0104 acta polytechnica ctu proceedings 40:104–110, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague hygro-mechanical model for concrete pavement with long-term drying analysis jakub veselý∗, vít šmilauer czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jakub.vesely.2@fsv.cvut.cz abstract. concrete pavements are subjected to the combination of moisture transport, heat transport and traffic loading. a hygro-mechanical 3d finite element model was created in oofem software in order to analyse the stress field and deformed shape from a long-term non-uniform drying. the model uses a staggered approach, solving moisture transfer weakly coupled with mps viscoelastic model for ageing concrete creep and shrinkage. moisture transport and mechanical sub-models are calibrated with lab experiments, long-term monitoring on d1 highway and data from 40 year old highway pavement. the slab geometry is 3.5 × 5.0 × 0.29 m, resting on elastic winkler-pasternak foundation. the validation covers autogenous and drying strain on the slab. the models predict drying-induced tensile stress up to 3.3 mpa, inducing additional loading on the slab, uncaptured by current design methods. keywords: concrete pavement, hygro-mechanical analysis, creep, shrinkage. 1. introduction concrete pavements present a proven solution for highways and airports due to its ability to withstand high mechanical loads with long service life compared to bituminous alternatives [1]. effects of wheel load, temperature and moisture generally need to be considered, as they influence stress, strain, deformations and service life [2]. for a detailed pavement analysis, finite element simulations with various factors have been carried out [3–7]. hydration and moisture transport generally induces autogeneous shrinkage strain, drying shrinkage strain with an impact on creep. early microcracks may occur due to internal or external restraints [8]. in addition, moisture gradients generate differential shrinkage between the top and the bottom, leading to warping [9, 10]. stress relaxation plays a role on the warping effect [11]. there are several approaches for creating a thermohygro-mechanical model for concrete. the most simple one uses a superposition principle, solving moisture, thermal, traffic load and asr effects separately [12]. due to a coupled nature of the problem, a staggered solution strategy offers more accurate analysis [13], solving moisture and heat transport and passing the obtained fields to a mechanical model. fully coupled problems with chemical submodels were proposed as well [14–16]. this paper presents a new hygro-mechanical 3d model, implemented in an open-source software oofem [17]. a staggered solution strategy uses solving of the weakly-coupled sub-models in the discretized time steps. the geometry uses 3d representation of a single rectangular concrete slab resting on winkler-pasternak foundation without dowel bars. the aim is to inspect overall performance of the concrete pavement under long-term drying loading scenario. a similar model has been published with limited calibration and interpretation [18]. 2. formulation of a hygro-mechanical model hygro-mechanical model solves transport of moisture and mechanical behavior of the slab. staggered solution strategy is adopted in the multiphysical model, partially separating transport transient problem from the mechanical one and passing transient moisture field to the mechanical problem. such an approach allows to define equivalent time in creep models or to calibrate constitutive laws independently. the models were implemented in oofem, an open-source and object-oriented software for finite element method [17]. 2.1. moisture transport a nonlinear moisture transport model describes concrete as a single-fluid medium with the governing equation: k(h) ∂h ∂t = ∇ · [c(h)∇h] , (1) where h is the dimensionless pore relative humidity, k(h) [kg/m3] is the humidity-dependent moisture capacity (k(h) = ∂w ∂h which is derivative of the moisture content w(h) [kg/m3] with respect to the relative humidity), c(h) [kg/m/s] is the moisture permeability in the bažant-najjar’s form: c(h) = c1  α0 + 1 − α0 1 + ( 1−h 1−hc )n   . (2) newton (convection) boundary condition is used in the form qn = β(h − h∞) with the moisture transfer coefficient β. 104 https://doi.org/10.14311/app.2023.40.0104 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 40/2023 hygro-mechanical model for concrete pavement e0 e1 e2 em η1 η2 ηm solidification cs s σ σ s εe εv εf εsh εt ε ksh αt figure 1. rheological scheme of mps model. 2.2. mechanical model the mechanical model consists of three different components. the first one takes a concrete slab with a viscoelastic model for concrete with ageing based on microprestress solidification theory (mps) [19]. the slab is placed on elastic, 2d subsoil, winklerpasternak model. the interaction between the slab and the subsoil is controlled by interface elements. 2.2.1. linear viscoelastic material model for creep and shrinkage a constitutive model for creep and shrinkage originate from b3 model [20] extended with mps formulation [19]. the rheological model in figure 1 shows non-aging asymptotic elastic spring, solidifying kelvin chain, aging dashpot, shrinkage strain, and thermal strain. the viscosity ηf in the flow strain εf depends on the evolution of the microprestress. all these units are connected in series, manifesting total strain decomposition into individual contributions. the governing equation for mps theory reads [21]: η̇f + 1 µst0 ∣∣∣∣ṫ ln h + ḣh ∣∣∣∣ (µsηf ) pp−1 = ψsq4 , (3) where µs is a parameter with the dimension of fluidity, p is a dimensionless material parameter influencing the size effect (for p = ∞, the size effect disappears), ψs is a temperature and humidity dependent factor, q4 is a material parameter, t is the current temperature, and t0 is the reference temperature. equation (3) can be simplified into equation (4) under the assumption of constant temperature t = t0 as: η̇f + k3 ∣∣∣∣ḣh ∣∣∣∣ ηp̃f = ψsq4 , (4) with parameters: p̃ = p p − 1 , (5) k3 = µ 1 p−1 s . (6) the drying shrinkage is computed as: ˙εsh,d = kshḣ, (7) with shrinkage ratio ksh. the autogenous shrinkage εsh,au is approximated as: εsh,au = ε∞sh,au [ 1 + (τau/te) w/c 0.38 ]rt,au , (8) where ε∞sh,au is the ultimate value of autogenous shrinkage strain, w/c is water/cement ratio and parameters τau,rt,au control strain evolution in time te. 2.2.2. winkler-pasternak subsoil model elastic subsoil is treated as a 2d winkler-pasternak model, capturing normal and shear stiffness with c1 and c2 parameters [22–24]. the governing equation reads: f(z) = c1w(z) − c2 ∂2w(z) ∂z2 , (9) where f(z) is surface load and w(z) is displacement. 2.2.3. interface elements the interface elements allow separation between the slab and subsoil, eliminating tension stress. both meshes of slab and subsoil share coinciding nodes at the interface. the traction-separation law takes the form: tn = knδ, kn = k for compression, (10) kn = 0.01k for tension, where δ is displacement between two nodes, positive in separation. shear stiffness is assumed zero. interface elements lead generally to slower converge or even convergence loss hence it is important to use reasonable time step to induce gradual slab deformations. 3. results and discussion the presented numerical model stems from a monitored highway slab. the pilot project started in 2017 as a joint activity among the road and motorway directorate (řsd čr), a contractor skanska, a.s. and the czech technical university in prague [25]. the project involved 8 978 m of concrete pavement built on d1 highway přerov-lipník nad bečvou, the czech republic. the pavement was cast between 07/2018 and 09/2019 with the opening to traffic dec 12, 2019. the binder used a slag-blended, slow hardening binder composed of 75 % cem i 42.5 r(sc) + 25 % ggbfs smš 400, corresponding to cem ii/b-s 42.5 n. 105 jakub veselý, vít šmilauer acta polytechnica ctu proceedings figure 2. vibrating wire gauges under protective covers. detail at one assembled location. 3.1. long-term monitoring system a long-term monitoring system was designed and installed in one concrete pavement slab with dimensions of 3.5 × 5.0 × 0.29 m. the system records temperature and strains at six measuring locations [25]. each location contains three vibrating wire strain gauges located 50 mm from the surfaces and in the mid-height. all the gauges have integrated temperature sensors. in addition, one thermal gauge was placed 150 mm under the pavement in order to deliver the sub-base temperature. ambient air temperature and solar radiation sensors were installed as well. two-lift concrete technology started with a bottom layer 240 mm thick with 370 kg/m3 of the binder, following with the top layer 50 mm thick with 420 kg/m3 of the binder. a two-step installation process was adopted, utilizing protective covers which hid the strain gauges before the casting, see figure 2. after the first finisher had passed, the covers were removed, the gauges put in their positions and the empty space filled back with concrete using hand vibrators, see figure 2. finally, the top layer finalized the slab. almost 55 000 m3 of concrete was placed in both layers, complying to required c30/37 strength class, see figure 3. vibrating strain gauges measure relative head displacements, which can be decomposed to: ε = εve + εt + εas + εds + εf + . . . , (11) where partial strains represent viscoelasticity, temperature effects, autogenous shrinkage, drying shrinkage, fracturing strain, etc. figure 4 shows partial strains on the mid-plane, zeroed at 2 hours after the end of setting for all gauges. autogenous shrinkage plays a dominant role in the first week, reaching −70 µε in the transversal direction (w62). in the longitudinal direction, continuous casting led to prestressing, which adds additional strain. a small average drying shrinkage strain from −30 to −120 µε is apparent after 3 years of drying. the measured strains allowed calculating the curvature of the slab, assuming a planar deformation of the cross-section. figure 5 shows the total curvatures, capturing the temperature variations and demonstrating a slow positive drift due to top drying. 0 10 20 30 40 50 60 0 10 20 30 40 50 60 c u b e m e a n c o m p r. s tr e n g th ( m p a ) concrete age (day) cem 𝖨 42.5 r(sc) mokrá, top layer bottom layer 75% cem 𝖨 42.5 r(sc) + 25% slag smš 400, top bottom 0 1 2 3 4 5 0 10 20 30 40 50 60 m e a n s p li t. t e n si le s tr e n g th ( m p a ) concrete age (day) cem 𝖨 42.5 r(sc) mokrá, top layer bottom layer 75% cem 𝖨 42.5 r(sc) + 25% slag smš 400, top bottom figure 3. compressive and tensile strength evolution of top and bottom concretes. -250 -200 -150 -100 -50 0 50 20 18 /0 8/ 01 20 19 /0 8/ 01 20 20 /0 8/ 01 20 21 /0 8/ 01 20 22 /0 8/ 01 p a rt ia l st ra in , ε -ε t ( 1 0 -6 ) location 1 location 2 location 3 (discarded) location 4 location 5 location 6 figure 4. partial strains on 6 locations. -0.5 0.0 0.5 1.0 20 18 /0 8/ 01 20 19 /0 8/ 01 20 20 /0 8/ 01 20 21 /0 8/ 01 20 22 /0 8/ 01 to ta l c u rv a tu re , κ ( k m -1 ) location 1 location 2 location 3 location 4 location 5 location 6 figure 5. total curvatures at 6 locations. 106 vol. 40/2023 hygro-mechanical model for concrete pavement parameter value unit parameter value unit m ec ha ni ca lm od el fc 45.0 mpa c1 70.0 mnm−3 cc cement content 370.0 kg/m 3 c2 60.0 mnm−1 w/c ratio 0.4 α 10.0 · 10−6 °c−1 a/c ratio 4.91 k 200.0 mnm−1 t0 7 days k3 35 τ1 3 · 10−5 days τm 5000 days ksh 1.5 · 10−3 e 35.0 gpa ε∞sh,au −87.0 · 10 −6 ν 0.2 τau 1.05 days rt,au −4.5 αau 1.166 m oi st ur e tr an sf er m od el c1 6.0 · 10 −4 kg/m·s α0 1.0 k(h) = k1 160.0 kg/m3 βtop 0.05 kg/m2 · d βbottom 0.01 kg/m2 · d table 1. list of material input parameters. 3.2. slab geometry the slab is meshed using 3d, quadratic, 20 node brick elements. interaction between neighboring slabs is neglected. in moisture transfer simulation, the influence of air and soil is represented with newton boundary conditions. the overview of used material parameters for concrete and subsoil is summarized in table 1. parameters ε∞sh,au,b4, τau, αau and rt,au control the evolution of autogenous shrinkage. the basic creep in the mps theory is influenced by the same four parameters q1–q4 as in the model b3, these are computed from fc, cc, w/c and a/c. the units of solidifying kelvin chain are defined with the lowest τ1 and the highest retardation time τm . time at the onset of drying is t0 = 7 days (for lab experiment t0 = 0.1 day), size effect for drying is neglected setting p = ∞. 3.3. hygro-mechanical simulation of long-term slab drying to simulate the long-term drying, the determination of material properties of concrete and relevant boundary conditions are necessary. material properties used in this paper are derived from drying shrinkage lab experiments. the kinetics of slab drying can be estimated from small lab samples 60 × 100 mm exposed to 50 % rh from the largest side. permeability (and hygric exchange coefficient) was calibrated from moisture transport model. boundary conditions are derived from a 40 year old concrete slab, which was removed during the demolition of d1 highway. the thickness of the old slab is only 240 mm compared to 290 mm used in current simulation. the moisture profile (see figure 6) and the fully saturated state were determined. we assume that the drying process is stabilized after 40 years, which identified ambient relative humidity as 65 % at the top and 80 % at the bottom. moisture transfer coefficients (βtop > βbottom) are set in such a way that the top drying becomes dominant. the linear moisture profile from figure 6 also implies that the permeability needs to stay constant, otherwise non-linear profile is obtained. the explanation for this phenomenon is likely a sorption isotherm and full resaturation with rain water. it was showed on prisms, that rewetting is orders of magnitude faster process than drying, thus occasional rain will lead to linear moisture profile [26]. this leads to setting α0 = 1 and c(h) = c1 according to equation (2). figure 7 shows shrinkage validation of numerical model both for lab experiment and for concrete pavement (strain gauge w52). the strain from gauge w52 was shifted by 40 µε to eliminate longitudinal prestressing. the predictions show that the slab will be drying for approximately 30 years. it should be mentioned that other gauges show smaller mid-plane shrinkage values, attributed likely to other partial strains and restraints. the initial drying takes place predominantly in top and bottom 100 mm of the slab, see figure 6. this leads to stress induction in these areas and subsequent redistribution in time. the early age stress field (first 2 years) shows tensile stress in the top part of the slab (around 3.3 mpa after 60 days and 2.2 mpa after 1 year, see figure 6). this will likely cause cracking in the slab when combined with temperature and traffic load. concrete slab should be able to cope with the resulting fracture process zone since the slab is continuously supported and controlled by displacement to a large extent. similar stresses were found on the analysis of restrained slab with thermo-hygro-mechanical model; exposing 150 mm thick slab to 60 % rh led to tensile stresses up to 3 mpa after 100 days [27]. 107 jakub veselý, vít šmilauer acta polytechnica ctu proceedings 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.5 0.6 0.7 0.8 0.9 1.0 d ep th ( m ) relative humidity (-) after 60 days after 1 year after 5 years after 15 years after 30 years real slab, after 40 years 0.00 0.05 0.10 0.15 0.20 0.25 0.30 -3 -2 -1 0 1 2 3 σx (mpa) figure 6. comparison of relative humidity profile (left) from fe simulation with real slab data after 40 years and σx (right) induced by long term drying. -800 -600 -400 -200 0 100 101 102 103 104 autogeneous shrinkageautogeneous shrinkageautogeneous shrinkage a u to g en ou s + d ry in g s tr ai n ( 1 0 -6 ) time of hydration (days) prism 100×60 mm, t0=0.1 d, rh=50% experiment at but slab 290 mm, t0=7 d, updated monitoring, w52+40 με figure 7. shrinkage validation of concrete slab with monitoring. after 2 years, the compression occurs within the top layer and the tensile stress occurs in the middle of the slab. the tensile stress reaches values up to 1.0 mpa, these are values lower than tensile strength of the concrete. the compression of the top layer of concrete is likely to help close the cracks developed in the earlier phase. the deformed shape induced by long-term drying is known as warping. the simulation predicts that the corners will raise up up to 1.7 mm when subjected to drying and self-weight (see figure 8). 4. conclusions this paper describes behavior of road concrete slab when subjected to the long-term drying. individual material models were calibrated based on lab and field experimental data, the main source being a long-term monitoring system on d1 highway. shrinkage was validated both for lab experiment and for the concrete pavement with moisture transport model. boundary conditions were calibrated from moisture profile of 40 year old pavement. drying induces high tensile stress on top surface, reaching 3.3 mpa after 60 days and 2.2 mpa after 1 year of drying. as the drying front advances, the stress profile changes and the surface becomes slightly compressed. slab warping is predicted as 1.7 mm corner uplift at the maximum. acknowledgements financial support for this work was provided by the czech science foundation under the project gačr 21-03118s and the czech technical university in prague, the grant sgs22/041/ohk1/1t/11. 108 vol. 40/2023 hygro-mechanical model for concrete pavement 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 100 101 102 103 104 c or n er u p lif t (m m ) time of hydration (days) concrete pavement subjected to drying figure 8. warping induced by long-term drying of concrete. references [1] m. g. lay. handbook of road technology. crc press, 2009. 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[26] l. dohnalová, p. havlásek, v. šmilauer. behavior of predried mature concrete beams subject to partial wetting and drying cycles. acta polytechnica ctu proceedings 34:1–5, 2022. https://doi.org/10.14311/app.2022.34.0001 [27] j. gomes, r. carvalho, c. sousa, et al. 3d numerical simulation of the cracking behaviour of a rc one-way slab under the combined effect of thermal, shrinkage and external loads. engineering structures 212:110493, 2020. https://doi.org/10.1016/j.engstruct.2020.110493 110 http://www.oofem.org https://doi.org/10.3390/ma13235530 https://doi.org/10.1016/j.cemconres.2014.03.008 https://doi.org/10.1007/bf02485965 https://doi.org/10.1016/j.cemconres.2016.04.001 https://doi.org/10.1016/j.trgeo.2021.100696 https://doi.org/10.1007/s11071-019-04977-9 https://doi.org/10.14311/app.2022.34.0001 https://doi.org/10.1016/j.engstruct.2020.110493 acta polytechnica ctu proceedings 40:104–110, 2023 1 introduction 2 formulation of a hygro-mechanical model 2.1 moisture transport 2.2 mechanical model 2.2.1 linear viscoelastic material model for creep and shrinkage 2.2.2 winkler-pasternak subsoil model 2.2.3 interface elements 3 results and discussion 3.1 long-term monitoring system 3.2 slab geometry 3.3 hygro-mechanical simulation of long-term slab drying 4 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0410 acta polytechnica ctu proceedings 38:410–416, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague exploring the state of knowledge and gaps regarding sustainability within viennese construction industry aída santana-sosa∗, martin aichholzer, elena mitrenova, martin schachenhofer vienna university of applied sciences campus wien – fh campus wien, department building and design, favoritenstraße 226, 1100 vienna, austria ∗ corresponding author: aida.santana_sosa@fh-campuswien.ac.at abstract. different approaches to assess sustainability goals are found in construction sector, mainly related to energy efficiency through building technology and high-tech components, to low-tech strategies based on passive planning methods, the use of renewable raw materials, and building certification programs, which, however, are mainly an evaluation of already built objects. austrian construction industry is extremely fragmented and heterogeneous, making research, development and innovation an issue for companies. the core of this project is to bridge the specific gaps in knowledge within viennese construction companies about sustainable architecture and respond to these urgent needs by developing a customized postgraduate course. within the first phase of the project, team meetings, a workshop and an online survey were run to assess the state of knowledge of viennese companies. by doing so, information deficits were identified, companies’ need for knowledge was determined and topics to be transferred were raised. keywords: sustainability, knowledge gap, expertise, industry needs, customized modular course. 1. introduction in 2016, construction industry in austria was responsible for 71.8 % of the total waste generated, being also the largest producer of co2 [1, 2]. strategies towards circular economy and the end of the fossil fuel era must be urgently implemented. use of renewable materials, improvement of thermal capacities of envelopes and reduction of operative carbon emissions among others, are imperative topics in the agenda of the city of vienna [3]. the assessment of such requirements is critical and need especial expertise of the companies involved. according to few surveys involving austrian construction companies [4, 5], the most relevant areas willing to be explored in this field are those related to low-tech approaches, energy efficient construction, ecological materials, and particularly life cycle assessment of buildings. based on that, it can be stated that there is a willingness of further qualification in sustainable construction among the professionals of the sector. besides, circular economy and building large-volume construction with renewable materials are relative ground-breaking issues, whereby targets and contents have not been widely explored, nor integrated in the curriculum of universities. for this reason the required expertise is rarely found within construction companies and seldom extended within the industry. main goal of this project is to bridge this gap between crosscutting issues in sustainable architecture and high education training, looking holistically at all sub-areas and serving like a platform for exchange, transfer and expansion of expertise, as well as for permanent networking and regional cooperation between companies, experts, students, alumni and those interested on sustainable planning. this project aims to extract, process, structure and offer specific know-how from different disciplines to professionals and interested public. for doing that, three phases are followed. the first phase consists on the analysis and representation of the current situation of the planning offices in vienna regarding the assessment of sustainability, their average existing level of knowledge and expertise and the specific topics of qualification willing to be explored. by doing so, knowledge gaps are identified. on this analysis relies the second phase, wherein customized public talks and interactive workshop series are developed, where the reported needs of qualification are intensively and interdisciplinary discussed with the purpose to identify enablers and barriers, and support the possible application of sustainable strategies. the third and last phase is focused on the further development of the findings reported in form of a post-gradual course, and the processing of collected knowledge to be offered to professionals and interested public within an open source platform. this paper describes the work done during the first phase, focused on the identification of knowledge gaps. 2. background sustainability aims to offer next generation at least same or higher quality of life than at current situation. it can be roughly divided into the pillars of economic, 410 https://doi.org/10.14311/app.2022.38.0410 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 exploring the state of knowledge and gaps regarding sustainability . . . ecological and socio-cultural quality. besides, technical, process and location quality play a major role in building sector. there are many different approaches to assess sustainability in terms of: (i) energy efficiency, mainly controlled by building technology and high-tech components; (ii) low-tech strategies, based on passive design concepts to maximize solar gaining and use of natural resources without high technological expenditure; (iii) renewable raw materials and timber construction; (iv) building certification programmes, which, however, mainly represent an evaluation of already planned objects; and (v) flexibility in floor plan design and participatory planning processes. a strongly arising aspect is the circular economy, which in construction means driving from the current linear system (cradle to grave) to a circular material and energy system (cradle to cradle), in which waste production, energy consumption and emissions are reduced by refusing, reducing and closing material and energy loops. austrian construction industry is scattered, small-scaled and heterogeneous, where 89 % of all austrian companies in the construction sector have 0–9 employees [6]. this situation makes research and development and thus innovation complicated. in addition, sustainable construction is currently struggling with the situation that every new building represents a prototype, meaning considerable additional work for planners, contractors and authorities. issues concerned to fire protection and sound insulation are rarely assessed with tested products. for this reason, several companies have already developed own individual solutions for partial and detailed problems. such in-house solutions could be shared, reducing complexity and increasing efficiency in design and construction. the existing certification and training courses found in austria are on the one side usually focused on specific and separated sub-areas like life cycle costs, straw bale construction, energy performance certificates and passive houses, or on the other side are extensive and expensive. workshops that can be booked individually covering different scales and areas of construction do not exist to date. main goal of this project therefore, is to work as knowledge hub to drive an exchange, transfer and development of know-how between viennese planning companies aiming to bundle the existing and dynamic expertise and to make it available to the public. 3. analysis of current situation in order to develop a customized curriculum and an adequate exchange platform adapted to the needs of the companies, an exhaustive analysis of the current situation was run. goal was to estate the average level of knowledge and knowledge requirements about sustainable planning within the sector. information deficits and learning needs were identified and key topics for deeper analysis and further exchange were uncovered. by doing so, the content of the curriculum and the exchange platform was assessed with the further willingness of collecting and discussing existing knowledge from different disciplines. besides, didactic methods and dedication intensity were also asked aiming to also meet the requirements of the industry, raising awareness and disseminating information on sustainable building. 3.1. data collection and target group aiming to assess correctly the industry needs from different disciplines, a three steps procedure was run when colleting the data: (i) collection and selection; (ii) discussion and development; and (iii) analysis and assessment. firstly, regular team meetings were run to collect ideas, best practices examples and measures for sustainable planning concepts, and select main topics and relevant actors. based on that, a professional workshop was organized where external experts on the defined topics were invited to enhance discussion, unfold sub-areas, fix boundaries and support further development. all inputs were documented and structured stating the basis for the formulation of an online survey, which settled the fundamentals of this study. its analysis and assessment determined the relevance of the selected topics and opened the repertory of new themes. primary target group for the data acquire was planning offices, evolving different stakeholders from spatial and landscape planners, to developers, architects, structural engineers and other specialist planners. secondary target groups were students, alumni, teachers and researchers in the field of sustainable planning. nevertheless other professional groups and construction trades were invited to take part implementing a bottom-up and a top-down approach. 3.1.1. collection and selection main objective of this step was to collect and filter existing knowledge and information on sustainable construction and select the essentials to make it available. literature and online review, together with regular team meetings and brainstorming served to define the key themes under the following discourse: sustainable architecture throughout the entire life cycle of a building englobes resource-related factors, socio-cultural issues and economic aspects. main strategy to ensure sustainability is the implementation of renewable, regional and less energy intensive resources, combined with a comprehensive use of technical equipment. on the socio-cultural level, the inclusion of all individual citizens is key for an active participation and co-determination to design responsible, resourcesaving and resilient cities. on the economic side, 411 a. santana-sosa, m. aichholzer, e. mitrenova, m. schachenhofer acta polytechnica ctu proceedings figure 1. definition of sustainability (grouped by rated values 1 to 2, 3 to 4 and 5 to 6, where 1 means not applicable at all and 6 fully applicable). affordability and circular economic system need to be assessed. 3.1.2. discussion and development main aim of this step was to work out the challenges and potential of sustainable buildings together with experts. an intensive workshop was held, where a total amount of 19 participants from various sectors of construction industry took part led by an experienced moderator. it also served as an official kick-off event for the public presentation of the project, where background, goals and methodology were introduced. the workshop was run in two rounds separated with a presentation of a project that served as best practices example in terms of ecological, economic and social sustainability. within the first round of the workshop, all participants were organized in groups of 4 or 5 according to their expertise and experience enhancing interdisciplinary exchange. groups were asked to discuss about their own concept of sustainable buildings and then asked to picture their developed concept, wherein the main essential characteristics were to be included and defined. each concept was presented and discuss with all participants, what generated and consolidated a common picture of a sustainable building. focus of the second round was the assessment of industry demands under the main concept of “needs and wants”. a brain-walk-method was performed, where four main topics were explored on four different flipcharts, englobing each thematic one to three key questions. all participants were encouraged to move freely through the panels and give answers and inputs to the themes prescribed. the intensive and interdisciplinary exchange performed served as a basis to the development of the online questionnaire, while settling the fundamentals of the upcoming curriculum. at the same time, concrete needs regarding further education and research were defined and existing information sources were identified. 3.1.3. analysis and assessment based on the discussion and inputs reported and documented by previous steps, an online questionnaire was developed following a 3-stage structure: • collection of personal data (profession, position, affiliation); • identification of the level of knowledge (own definition of sustainability and challenges when implementing sustainable strategies); and • analysis of further training demands (knowledge gap, needs and prerequisites). email addresses and linkedin profiles from potential interested professionals were selected and grouped. the survey was sent to a total amount of 15 500 contacts. by doing so, the project was also promoted within the sector, arising the awareness of the thematic. 3.2. findings 423 people were reached and a total of 121 participants took part in the survey. the results are representative with the respective range of fluctuation within a confidence interval of 95.5 %. 3.2.1. participants’ professional background almost half of the participants taking part on the questionnaire are architects (43.8 %), being consequently the most representative group, followed by researchers and lecturers (20.6 %), and advisers (19.0 %). further professionals took part in the survey representatively such as developers (16.5 %), specialist planers (12.4 %) and decision makers (9.1 %). a deeper analysis regarding to specific job positions showed that half of respondents are managing directors (48.6 %), while one fourth employees (27.9 %) and team leaders (23.4 %) respectively. on the question of how often participants deal with sustainability in their daily activities, 40 % said very often to often, while 20 % said little or very little, and another 40 % said only sometimes. these data represent a higher participation of stakeholders involved in the process of decision-making, what brings relevance to the survey and the project as they can work as enablers having relevant impact on further steps in production and construction processes. 3.2.2. definition of sustainability and barriers to its implementation for the definition of the multifaceted concept of sustainability in construction, almost two thirds of the respondents answered that a holistic approach with a responsible use of resources and an intensification of reuse, renovation and re-densification is fully applicable with an average of 5.38, 5.37 and 5.31 respectively, where 1 means not applicable at all and 6 fully applicable. in figure 1 these most suitable definitions 412 vol. 38/2022 exploring the state of knowledge and gaps regarding sustainability . . . figure 2. ability to implement sustainable strategies. of sustainability are represented, where the 86.7 % of all respondents scored a holistic approach with the highest values 5 and 6, followed by the responsible use of resources with 85.6 %, and renovation or restoration with 84.4 %. by these three definitions just around 4.4 % up to 10.0 % of all participants scored them with the lower values of 1 and 2, and 3 and 4. in further stage, the optimisation of buildings physics with an average of 4.91 was scored by the 71.1 % with the highest values 5 and 6, by the 21.1 % with 3 and 4, and by the 7.8 % with 1 and 2. next, the use of renewable energy towards zero-energy and plus-energy got a value of 4.82, where the 72.7 % of participants rated them with 5 and 6, 18.2 % with 3 and 4, and 9.1 % with 1 and 2. last, participation in design process was also highlighted as key aspects with an average score of 4.40, being valued the highest by the 51.1 %, with 3 and 4 by the 40.0 %, and with 1 and 2 by the 8.9 % of respondents. the current situation regarding the capacity or ability of implementing sustainable strategies is rated just as sufficient by the participants, who stated an average of 3.7 on a scale from 1 to 6, where 1 means they cannot implement nothing at all and 6, everything. that means, more than half of participant reported that they are able to only implement an enough amount of strategies, while the other half meant many and very few strategies almost equally. figure 2 represents the quoted answers. main reasons quoted as barriers for a successful implementation of sustainable strategies in design and construction were the lack of: (1.) awareness of the client (5.18 to 6); (2.) willingness to invest time and personnel resources (4.75 to 6); (3.) tax approaches and subsidies (4.72 to 6); (4.) marketing or public information (4.51 to 6); (5.) know-how of the executing companies (4.47 to 6); (6.) cost certainty (4.25 to 6); (7.) expertise of the planners (4.20 to 6); (8.) regulatory or normative basis (4.15 to 6); (9.) willingness to cooperate and share knowledge (4.10 to 6); (10.) specific education and training (3.87 to 6). access to specific knowledge and training opportunities play a comparatively small role, in contrast to the lack of awareness of the client. therein lies the need to provide suitable accompanying measures, which depending of target groups are to be offered in different forms. in order to arise awareness among clients, dissemination, press work or events could be promote its consideration. regarding to architects, trainings for effective advisory talks could be developed, based on conducting discussions, with argumentative discourses emphasising objective advantages and personal benefits, together with opposite foundations to debate, with the aim to bring out decision and generate positive conclusions. 3.2.3. information sources the sources of information consulted by the participants when aiming to design and build more sustainably are mainly specific literature and reports (70.2 %) and personal contacts (67.0 %). second main information sources are further education formats and trainings like seminars or workshops (47.8 %), conferences, messes, and congresses (45,7 %), and online platforms (45.7 %), followed closely by contacting directly planners with experience and representative projects (43.6 %). certification tools, guidelines, building regulations and standards were rated just by 33.0 % and 29.8 % respectively. consulting personal contacts can be considered ambivalent since real sources of knowledge like certification tools, guidelines, etc. as well as building regulations and standards are not extensively used and therefore the level of knowledge of consulted contacts cannot be stated. 3.2.4. knowledge gap and demand by analysing the knowledge gaps, the life cycle assessment of constructions as shown in figure 3 was reported by 59.3 % of the participants as the area with highest score, wherein the 19.8 % and 20.9 % considered there is enough to much available knowledge. since the lifecycle assessment under a holistic approach was also quoted as an accurate definition of sustainability, it should be considered a fundamental aspect when developing a curriculum towards sustainable construction. planning optimized life-cycle costs buildings was also considered by the 43.0 % of the respondents as an area with remarkable knowledge gap, wherein the 38.4 % stated there is enough knowledge, 413 a. santana-sosa, m. aichholzer, e. mitrenova, m. schachenhofer acta polytechnica ctu proceedings figure 3. knowledge gaps (grouped by rated values 1 to 2, 3 to 4 and 5 to 6, where 1 means high gap of knowledge and 6 no gap at all). figure 4. demand of knowledge (quoted from 1 to 6, where 1 means not applicable at all and 6 fully applicable, and starting by 3.00). and 18.6 % declared there are no relevant knowledge gaps. these statements make clear that a compromise between ecological and financial assessment need to be widely explored taking into account the whole lifecycle. specific expertise on sustainable planning approaches and confidence on using tools and software for this purpose were reported by 32.5 % and 35.3 % respectively as aspects with high knowledge gaps willing to be assessed. however, 43.0 % and 50.6 % of the respondents considered that there is enough expertise, where the 24.4 % and 14.1 % respectively considered there are no need for further knowledge building in this regard. these two aspects are strongly related with the two previous ones, since they englobe the required competences and tools for achieving the desired lifecycle assessment rated with the highest scores. finally, building services, building physics and structural performances were pointed out by 27.1 %, 24.4 % and 16.7 % each, meaning that even though being significant areas to consider when planning sustainably, enough expertise is already extended within the sector. that becomes evident with the 52.9 %, 38.4 % and 51.2 % which considered there is enough expertise, and with the 20.0 %, 37.2 % and 32.1 %, which stated there is no knowledge gap. these last aspects can also be grouped into a category which refers to the performance of the building. by doing so, three main groups are recognised where all key knowledge gaps can be embedded. the first and more relevant refers to the correct and accurate planning regarding the whole lifecycle of the building including its ecological and economic assessment. the second group relates to the required expertise and tools to be able to achieve those in the design. the third describes the structural, acoustic and thermal performance of a construction and its correlated building services. by assessing the specific project phases, wherein to focus the implementation of strategies, all phases were rated as important, from basic analysis, to preliminary design, design planning, submission planning, detailed design, tendering, supervision of construction work and documentation, and project management, wherein basic analysis and submission planning was respectively rated with the maximal value (5.18) and the minimal one (4.12) with little difference between. that means all contents are above average in importance, having received the highest value for answer 6 (from 55 % to 28 %), except for submission planning, which highest percentage at answer was 4, what is still above average. the survey also highlighted areas represented in figure 4 as key topics for desired further training, where almost all contents are important to an above-average extent, but without peak. the top demand is formed by a sextet with values between 4.94 and 4.48, namely: (i) deconstruction and disposal, englobing urban mining and dismantling and reuse and recycling of building components (4.98); (ii) life-cycle management including economic and ecological assessment (4.78); (iii) low-tech planning strategies and climate-friendly 414 vol. 38/2022 exploring the state of knowledge and gaps regarding sustainability . . . figure 5. effective strategies (quoted from 1 to 6, where 1 means not applicable at all and 6 fully applicable, and starting by 4.00). architectural concepts (4.75); (iv) building with renewable materials, especially large volume timber constructions, assessing integral planning, structural performance, technical building services and building physics (4.71); (v) sustainable renovation, revitalisation and redensification of existing buildings (4.65); and (vi) innovative materials and developments (4.48). these data reflect a clear trend towards passive and responsible architectural design with low-tech planning strategies that use innovative environmentally friendly, renewable and low-carbon materials, addressing the full life cycle of the building and encouraging the extension of its lifespan through renovation, revitalisation or transformation (change of use) and re-densification, as well as through deconstruction and recycling of its building elements and components. on the other side, the respondents rated with the relatively lowest value, the need for theory and history (3.15), the use of certification tools and systems (3.71), and the social participation in planning processes (3.86). the low rate of the first aspect could be justified through the presumed remoteness from practice and questionable applicability and practicability. less comprehensible is the low interest in certification systems, which actually ease remarkable improvements within the overall process. these certification tools and systems are normally used after the project is planned and built in order to gain recognition, so that they do not impact the design development. with this in mind, it could be of relevant importance to implement these certification systems and tools proactively in order to manage different solutions within design and quantify their repercussions, helping in the decision-making process. low rated were also social aspects like participation in planning processes, what does not fit with one of the main pillars of sustainability, namely the social sustainability, and is an issue that need to be assessed, having architects the chance to work as advisors to clients and society communicating benefits of building responsibly and arising the awareness regarding sustainability. in between we find a number of topics that were scored between 3.96 and 4.33, which englobe strongly related aspects like design concepts towards passive house, building physics and energy concepts, and software tools and simulation. other aspects found within these scores relate to the development of urban and rural areas, and to promotional approaches and fiscal policies. the low scores for these aspects could be justified by considering that they are competencies that are beyond the reach of the planning offices, which are the group with the highest participation in the survey. regarding the methodological design of the customized programme was answered by the majority with the preference of part-time continuing education courses as evening or weekend programmes (67.1 %), as well as the possibility of the completion of individual modules (59.5 %). getting a certificate (44.0 %) was still rated as considerable relevant, while an academic degree (11.4 %) as low important. undesirable was the fact of completing a large course of about 2 years (6.3 %). 3.2.5. possible effective strategies all given strategies were considered by the participants to be more effective than average with a punctuation between 4.28 and 5.50. nevertheless, as shown in figure 5, a top quintet can be recognized, formed by: (i) raising awareness for sustainable building (5.50); (ii) networking and cooperation (5.22); (iii) tax approaches and subsidies (5.16); (iv) easy access to practical solutions (5.12); and (v) (re)use of existing buildings (5.00) instead of new construction with average values between 5.50 and 5.00 of 6.00. three of those topics, namely networking and cooperation, easy access to practical solutions and (re)use of existing buildings, are central concerns of this project, while the fact of raising client’s awareness about sustainability was mentioned in the previous section as important topic to be addressed despite being scored 415 a. santana-sosa, m. aichholzer, e. mitrenova, m. schachenhofer acta polytechnica ctu proceedings low, eventually through architects acting as advisor. the other theme requires a rethink about the inclusion of other stakeholders, especially politicians with regard to tax approaches and subsidies. besides, owners and investors could be also key stakeholders to be included as target group in order to explore the possibilities towards (re)use of existing buildings rather than new construction. on the other side, even the three strategies considered as least effective, namely: (i) promotion of participation, co-determination and other democratic strategies (4.28); (ii) transfer of self-generated knowledge (4.30); and (iii) mandatory monitoring of the entire life-cycle of buildings (4.37), were rated by almost 50 % with 5 and 6 (average values around 4.30), meaning that they are also considered of highly relevance. more transparency eventually through public benchmarks (4.99) was also stated as remarkably important, followed by a clear and common definition of sustainable construction (4.87), alternative forms of financing (4.72), by regulatory or normative obligations (4.71) and basic business knowledge (4.62). 4. conclusions and next steps having explored and analyzed the knowledge gaps, demands and desired approaches of the planning offices during the first phase of the project, a series of interdisciplinary and interactive workshops will be organized within the second stage of the project with the aim of fostering tailor-made and diversity-adapted mutual knowledge transfer for design and planning firms. the main objective of this exchange of different viewpoints and professional approaches is to force a fruitful discussion in order to identify facilitating and constraining aspects of the implementation of the proposed strategies. in this way, the topics chosen for a more detailed analysis are grouped and will be discussed in thematic blocks. experts on each topic will be invited to take part not only in the presentation of the contents, but also to actively interact in the discussion with the participants. in the same way, a building already realized and recognized for its sustainable quality will be chosen and gradually analyzed according to the topics relate to each thematic block. in this way, the participants together with the invited experts and the planners and builders of the building themselves will evaluate the implemented strategies by applying the acquired knowledge. parallel to the analysis, new strategies will be proposed and their applicability will be jointly evaluated. besides these interactive workshops, and according to the analysis run within the first stage of project, public lectures and expert talks are to be organized in order to raise general awareness and disseminate information about sustainability to interested public. both methods of knowledge transfer will be evaluated through questionnaire and discussions to form the basis for the development of the customized postgraduate modules within the third stage of the project, which will be incorporated in the existing university training format. furthermore, all knowledge gained throughout the whole project will be processed and offered to the public, so that they can serve as a basis for new services, research areas and product developments. in further steps within the third stage of the project, an extended literature review on educational theory and construction management will be performed to propose a suitable and functioning learning framework, where concepts like competence will be defined including its types, orders and domains of learning [7], to better address the didactic methods and contribute to the sector. acknowledgements authors want to thank everyone involved in this project, from those who took part into the online survey, to the experts working in the workshops. a special thank goes to the founder agency ma23 for enhancing sustainability on construction and education. references [1] bundesministerium für nachhaltigkeit und tourismus. die bestandsaufnahme der abfallwirtschaft in österreich. (statusbericht 2018). bundesministerium für nachhaltigkeit und tourismus, wien, 2018. [2] international energy agency, united nations environment programme. 2018 global status report: towards a zero-emission, efficient and resilient buildings and construction sector, 2018. [2022-04-06]. https://wedocs.unep.org/20.500.11822/27140 [3] magistrat der stadt wien. klimaschutzprogramm der stadt wien: fortschreibung 2010-2020, 2009. [4] kompetenzzentrum bauforschung. mitgliederbefragung österreichisches baugewerbe: spezialauswertung wien, 2017, 2017. [2022-04-06]. https://www.zukunft-bau.at/sites/default/files/ dateien/inhalt/180227_folder-umfrage_wien.pdf [5] s. wirth. die größten herausforderungen werden noch wichtiger, 2016. in: expertenbefragung zukunft bauen, [2022-04-06]. https://www.expertenbefragung.com/index.php/ zukunft-bauen/zukunft-bauen-2016/groessteherausforderungen-werden-noch-wichtiger [6] statistik austria. bundesanstalt statistik. ergebnisse im überblick: statistik zur unternehmensdemografie 2007 bis 2019 nach wirtschaftsbereichen, 2017. [2022-04-06]. https://www.statistik.at/fileadmin/ pages/178/1_ergebnisse_im_ueberblick_statistik_ zur_unternehmensdemografie_2007_bis_2020.ods [7] g. killip. a reform agenda for uk construction education and practice. buildings and cities 1(1):525–537, 2020. https://doi.org/10.5334/bc.43 416 https://wedocs.unep.org/20.500.11822/27140 https://www.zukunft-bau.at/sites/default/files/dateien/inhalt/180227_folder-umfrage_wien.pdf https://www.zukunft-bau.at/sites/default/files/dateien/inhalt/180227_folder-umfrage_wien.pdf https://www.expertenbefragung.com/index.php/zukunft-bauen/zukunft-bauen-2016/groesste-herausforderungen-werden-noch-wichtiger https://www.expertenbefragung.com/index.php/zukunft-bauen/zukunft-bauen-2016/groesste-herausforderungen-werden-noch-wichtiger https://www.expertenbefragung.com/index.php/zukunft-bauen/zukunft-bauen-2016/groesste-herausforderungen-werden-noch-wichtiger https://www.statistik.at/fileadmin/pages/178/1_ergebnisse_im_ueberblick_statistik_zur_unternehmensdemografie_2007_bis_2020.ods https://www.statistik.at/fileadmin/pages/178/1_ergebnisse_im_ueberblick_statistik_zur_unternehmensdemografie_2007_bis_2020.ods https://www.statistik.at/fileadmin/pages/178/1_ergebnisse_im_ueberblick_statistik_zur_unternehmensdemografie_2007_bis_2020.ods https://doi.org/10.5334/bc.43 acta polytechnica ctu proceedings 38:410–416, 2022 1 introduction 2 background 3 analysis of current situation 3.1 data collection and target group 3.1.1 collection and selection 3.1.2 discussion and development 3.1.3 analysis and assessment 3.2 findings 3.2.1 participants' professional background 3.2.2 definition of sustainability and barriers to its implementation 3.2.3 information sources 3.2.4 knowledge gap and demand 3.2.5 possible effective strategies 4 conclusions and next steps acknowledgements references my title https://doi.org/10.14311/app.2022.36.0206 acta polytechnica ctu proceedings 36:206–215, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague probabilistic models for resistance variables in fib model code 2020 for design and assessment miroslav sýkoraa, ∗, wouter botteb, robby caspeeleb, dimitris diamantidisc, aurelio muttonid, raphaël d. j. m. steenbergenb, e a czech technical university in prague, klokner institute, šolínova 7, 166 08 prague, czech republic b ghent university, faculty of engineering and architecture, department of structural engineering and building materials, technologiepark-zwijnaarde 60, 9052 zwijnaarde, belgium c oth regensburg, faculty of civil engineering, galgenbergstrasse 30, 93053 regensburg, germany d école polytechnique fédérale de lausanne, school of architecture, civil and environmental engineering, station 18, ch-1015 lausanne, switzerland e tno delft, department of structural reliability, p.o. box 155, nl-2600 ad delft, netherlands ∗ corresponding author: miroslav.sykora@cvut.cz abstract. the fib model code offers pre-normative guidance based on the synthesis of international research, industry and engineering expertise. its new edition (draft mc 2020) will bring together coherent knowledge and experience for both the design of new concrete structures and the assessment of existing concrete structures. this contribution presents an overview of the main developments related to the partial factors for materials. in the draft mc2020, the partial factors are presented in tables for clusters of cases depending on consequence classes and variability of basic variables. furthermore, formulas and background information are provided to facilitate updating of the partial factors. this contribution discusses the different assumptions adopted in mc 2020 for design and assessment. main changes with respect to the previous version are related to description of the difference between in-situ concrete strength and the material strength measured on control specimens, and to modelling of geometrical variables. the presented comparison of the requirements imposed by eurocodes and mc 2020 for design reveals insignificant differences. the assessment requirements may be decreased by about 25% when the conditions specified in mc 2020 are satisfied. hence, the revised mc 2020 will provide designers and code makers with wider possibilities to utilise actual data and long-term experience in assessments of existing structures. keywords: assessment of existing concrete structures, eurocodes, fib model code 2020, partial factors, probabilistic models, reliability, updating. 1. introduction fib (international federation for structural concrete) has been systematically revising its flagship document fib model code (mc). the mc offers pre-normative guidance, synthesis of international research with industry and engineering expertise, and advanced tools for international code writers as well as industry practitioners. the main aspiration of its new edition (draft mc 2020) is to bring together coherent knowledge and experience for both design and assessment and to provide a single code for both new and existing concrete structures [1]. the fib mc serves as a pre-standard, intended to provide the basis for development of future eurocodes and other international standards. under fib commission 3 existing concrete structures, task group 3.1 is drafting the mc 2020 sections on: • reliability requirements (target reliability levels for various limit states in design and assessment and different consequence classes) and • the partial factor method for design and assessment. this contribution presents an overview of the main developments related to the latter topic, focusing on resistances. it is demonstrated that the revised mc 2020 should provide designers and code makers with wider possibilities to utilise actual data and longterm experience in the assessment of existing concrete structures. 2. probabilistic models for resistances 2.1. general the draft mc 2020 section on the partial factor method is focused on the analysis of basic structural elements such as beams, slabs and columns through simplified (analytical) models. the safety formats 206 https://doi.org/10.14311/app.2022.36.0206 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 36/2022 probability of fib model code 2020 resistance variables fc on control specimens vfc = 0.1 conversion factor for in-situ cast concrete (μη = 0.95, vη = 0.12) concrete section area (μa ≈ 1 and va ≈ 0.04) design resistance model uncertainty – column crushing (μθr = 1.02, vθr = 0.07) fc in-situ vfc,is = 0.08 or 0.15fc in-situ vfc,is = 0.08 or 0.15 concrete section area (μa ≈ 1 and va ≈ 0.01) concrete section area (μa ≈ 1 and va ≈ 0.01) assessmentassessment resistance model uncertainty – column crushing (μθr = 1.02, vθr = 0.07) figure 1. basic variables considered in resistance model for short columns in design and assessment situations. to be applied in conjunction with advance numerical models are also addressed by mc 2020 but they remain beyond the scope of this contribution. mc 2020 emphasises the importance of updating of the models applied in reliability verifications considering all available information about the structure. however, to assist routine applications, the model code also provides the conventional probabilistic models for parameters of resistances and load effects considered to be applicable for common reinforced concrete (rc) structures with a reasonable level of approximation. these conventional models are mostly based on the probabilistic models assumed in the background document to the revised pren 1992-1-1:2021 for design and assessment of concrete structures [3] and they are consistent with those provided by the jcss probabilistic model code [4], and with the recommendations of fib bulletin 80 [5]. focusing on non-deteriorated rc structures, the flexural resistance of members and the resistance of short columns under compression, r, are obtained as the product of a resistance model uncertainty, θr, geometrical property, a, and of material strength, f : r ≈ constant × θr × a × f (1) where the strength f can cover the factor, η, accounting mainly for the difference between the compressive concrete strength measured on control specimens (fc) and the in-situ strength (fc,is), and then fc,is ≈ η × fc. the resistance r is assumed to be lognormally distributed. it is assumed that reinforcement properties are governing bending failures whereas concrete compressive strength dominates in compressive crushing of columns without significant second order effects. figure 1 shows the basic variables considered in the resistance model for short columns in design and assessment situations. the flexural resistance, dominated by reinforcement yielding, is described in a similar way. table 1 provides the probabilistic models for resistance parameters considered in mc 2020. µ denotes the bias of a variable systematic deviation of random values of the variable from its characteristic (nominal) value, considered as the ratio of mean to characteristic (nominal) value; v is a coefficient of variation. 2.2. material properties the models for concrete strengths, fc and fc, is, and the conversion factor, η, provided in table 1 are related to ordinary strength concretes (cylinder concrete strengths approximately up to 100 mpa) and standard quality related to in-situ cast rc structures. for high-strength or precast concretes or other than standard execution quality, other models may apply [6]. the coefficient of variation (cov) of concrete strength commonly decreases with increasing mean strength; see also [7]. for instance, torrenti and dehn [8] proposed for in-situ cast concrete: vf c ≈ 0.1 × (fc/40 mpa)3/2 (2) the value of µη in table 1 represents concretes in tops of columns; for other governing regions the bias may be increased to 1.0 or up to 1.05 for bottoms of columns. similarly, for smaller members (h < 450 mm) lower biases µη = 0.95 may be considered while 207 m. sýkora, w. botte, r. caspeele et al. acta polytechnica ctu proceedings basic variable x situation dist. bias µx coeff. of variation vx note concrete compressive strength measured on control specimens (in-situ cast concrete), fc design ln exp(1.645vf c)= 1.18 0.1 µf c = 1.10 and vf c = 0.06 could be considered for precast concrete. concrete compressive in-situ strength, fc, is assessment ln exp(1.645vf c,is)= 1.14 − 1.28 0.08− 0.15 to be estimated from core tests. conversion factor for in-situ cast concrete, η design ln 0.95 0.12 based on [2]. for precast concrete, lower coefficient of variation can apply. yield strength of steel reinforcement, fy design and assessment ln [exp(1.645vf y )] = 1.08 0.045 no conversion factor is commonly applied (µη = 1 and vθ = 0). concrete section area, ac design n 1 0.04 the case of normal uncertainty in geometry considered as the default case, width of the column ∼ 300 mm. mc 2020 provides guidance for other cases. assessment n 1 0.015 should be based on measurements in the existing structure. normal uncertainty in geometry as adopted for design may be considered in the absence of measurements. see section 3 for details. effective depth, d design n 1 0.04 normal uncertainty in geometry as the default case, effective depth ∼ 200 mm. mc 2020 provides guidance for other cases. assessment n 1 0.01 to be based on measurements in the existing structure. normal uncertainty may be considered in the absence of measurements. flexural resistance model uncertainty, θr design and assessment ln 1.09 0.045 considered to derive γs in both design and assessment situations. resistance model uncertainty for crushing of columns without significant 2nd order effects, θr ln 1.02 0.07 considered to derive γc in both design and assessment situations. table 1. probabilistic models for resistance parameters considered in draft mc 2020. 208 vol. 36/2022 probability of fib model code 2020 resistance variables for larger members (h ≥ 450 mm), more favourable values may be applied (e.g. µη = 1.03). mc 2020 indicates that the conversion factor can be influenced by the effects of non-standard quality of execution, various types of binding materials, seasonal aspects, size of the structure etc. more details including an extensive literature review and the analysis of available test results can be found in the background document to pren 1992-1-1 [3]. note that the conversion factor η is implicitly covered by test results for in-situ compressive strength, fc, is, and µη = 1 and vη = 0 are considered when the model for fc, is is based on core tests. 2.3. geometrical variables regarding geometrical variables, it is assumed that the section area, ac, relates to the failure modes governed by concrete crushing (typically columns under compression) while the effective depth, d, should be taken into account when reinforcement yielding is dominating (typically bending of beams or slabs). the case of normal uncertainty in geometry is considered as the default case for design. for the section area of a square column with fully correlated dimensions b × b, no bias is assumed and the cov of ac can be obtained as: vac = 2vb = 0.04 × (300 mm/b)2/3 (3) when the two widths are statistically independent, the factor of 2 should be replaced by √ 2 and thus: vac = 0.0283 × (300 mm/b)2/3 (4) in the case of doubts about a level of correlation, use of equation 3 is recommended. when considering normal uncertainty in geometr y, the statistical characteristics of effective depth can be considered as a function of its nominal value, dnom: µd = 1 − 0.05 (200 mm/dnom)2/3 vd = 0.05 (200 mm/dnom)2/3 (5) it is further considered that uncertainty in geometry may be reduced when a decisive geometrical property is measured in the existing (finished) structure. this is considered as the default case for the assessment of existing structures. µac = 1 and vac = 0.015 may then be considered for the section area, and µd = 1 and vd = 0.01 for the effective depth. further, uncertainty in geometry can also be reduced when an increased execution quality and quality control are required. for instance, when higher execution quality is reached (ensuring that the geometrical deviations of tolerance class 2 according to en 13670 on execution of concrete structures are fulfilled), µac = 1 and vac = 0.02 may be considered for the section area. for the effective depth, the following characteristics may be considered: µd = 1 − 0.025 (200 mm/d)2/3 vd = 0.025 (200 mm/d)2/3 (6) it can be shown that the partial factor for concrete, γc , is insignificantly affected by the variability of the section area in common cases. in contrast, covs of the effective depth, yield strength and model uncertainty for bending are of a similar magnitude and the variability of the effective depth should be adequately considered when determining the partial factor for reinforcement, γs . considering a reference period of 50-years, a recommended value of the sensitivity factor for resistance is αr = 0.8 and the target reliability index for design and consequence class 2 (medium failure consequences) is β = 3.8 in mc 2020. assuming that the three resistance variables (fy , d, θr) have a similar effect on the design or assessment value of the flexural resistance, the sensitivity factor for the effective depth is estimated as αd = 0.8/ √ 3 = 0.46 and its design value should approximately correspond to a 4% fractile, φ(0.46 × 3.8) = 4 % (with φ denoting the cumulative distribution function of the standardised normal distribution). for assessment situations, target reliabilities lower than those for design are considered in mc 2020 and a slightly higher fractile would characterise the assessment value of the effective depth. to provide indications for both situations, figure 2 displays 5% fractiles of the effective depth as a function of its nominal value, dnom (the fractiles are normalised with respect to dnom). the models for effective depth described in this section are applied. figure 2. effective depth 5% fractiles as a function of nominal value. figure 2 shows that the probabilistic models recommended for the effective depth in mc 2020 (for d = 200 mm) yield rather lower estimates for the cases of improved quality and particularly for normal 209 m. sýkora, w. botte, r. caspeele et al. acta polytechnica ctu proceedings uncertainty that are primarily relevant for design situations. in the latter case for dnom = 1000 mm, the design value of effective depth increases by about 8% in comparison to the reference case based on dnom = 200 mm. in contrast, measurements in the existing structure are associated with the smallest uncertainty and the highest design or assessment value of effective depth is obtained. regarding considerations of the variability of effective depth in design, mc 2020 provides an additional simplification. the variability of d may be ignored (µd = 1 and vd = 0) if the design resistance is based on the design value of effective depth obtained as dd = dnom − ∆d where: • dnom is determined on the basis of the nominal cover. • ∆d is the deviation of the effective depth with: ▷ ∆d = 15 mm for reinforcing and post-tensioning steel, ▷ ∆d = 5 mm for pre-tensioning steel 2.4. model uncertainties in the background document to pren 1992-1-1 [3], the statistical characteristics for resistance model uncertainties adopted in mc 2020 (table 1) are based on the comparison of test and model results. the model uncertainty is estimated as the ratio of test resistance to a model estimate [9]. for short columns with negligible second order effects, moccia et al. [6] investigated cases without and with eccentricity for cylinder concrete strengths up to 100 mpa. they found a bias of 1.02 and cov of 0.087. the latter is affected by the test variability and the uncertainties in the test parameters. considering typical variabilities for these effects, the vθr-value reduced to 0.07. the statistical characteristics for flexural resistance, µθ r = 1.09 and vθr = 0.045, were derived in the background document [2] in a similar way. bending tests were considered for various reinforcement steel classes and calculated strains around 1-1.5%. both bias and cov tend to increase with increasing calculated strain. in the draft mc 2020, the resistance model uncertainties are assumed to have a lognormal distribution [4]; for further information see also annex a of fib bulletin 80 [5] or the numerical study by sykora et al. [10]. note that slightly less favourable resistance model uncertainty characteristics (bias around unity and higher cov) are indicated in mc 2020 complex situations analysed by advanced numerical models [11]. 3. updating of probabilistic models when specific information about the structure under investigation is available, the probabilistic models for materials and actions and subsequently related partial factors can be updated. particularly in the assessment of an existing structure, uncertainties in resistances and load effects can often be reduced on the basis of inspections, measurements, and tests. the experience from practical applications suggests that it is often highly beneficial when in-situ concrete strength is investigated and hence the important uncertainty in the conversion factor is eliminated. however, it is also possible that uncertainties to be considered in the reliability verification of the existing structure exceed those considered for a relevant design situation and the partial factors adopted for design might be insufficient. the prospective eurocode on the reliability assessment of existing structures, pren 1990-2:2021, makes distinction between the preliminary and detailed assessment; the latter being in the main focus of this contribution. regarding the updating of basic variables, the preliminary assessment is typically conducted assuming the default characteristics of basic variables similar to those for new structures while these characteristics are updated for the detailed assessment or in the case of doubts about an appropriate model for the basic variable. if justified to be relevant for the structure under investigation, prior information can be combined with new information obtained e.g. by measurements and/or tests through bayesian updating. mc 2020 provides no information about the strength of the prior information provided by the presented conventional models. following the background document to the reliability basis in eurocodes [12], it might be assumed that the mc 2020 resistance models are based on prior information associated with the equivalent sample sizes n′ = 1 to 5 for the mean values (biases) and v′ = 3 to 10 for standard deviations. typically, lower values apply for concrete and higher for steel reinforcement. more details can be found in the jcss probabilistic model code [4] and the jcss monograph on assessment of existing structures [13]. 4. recommended values of partial factors in draft mc 2020 the draft mc 2020 elaborates on a partial factor method for the reliability assessment of new or existing structures conforming to iso 2394:2015 for structural reliability principles and en 1990:2004 for basis of structural design. 4.1. partial factors for design for design with respect to ultimate limit states (ulss), the partial factor for materials is obtained according to mc 2020 as: γm = rk/rd ≈ [exp(−1.645 vf )]/[ µθr µa µη exp(−αr β √ v 2θr + v 2a + v 2η + v 2 f ] (7) 210 vol. 36/2022 probability of fib model code 2020 resistance variables cc γc γs γg self-weight γg other permanent loads γq (imposed) γq (wind) cc1 1.4 1.1 1.2 1.3 1.3 1.6 cc2 1.5 1.15 1.25 1.35 1.5 1.85 cc3 1.6 1.175 1.25 1.4 1.7 2.1 table 2. recommended values of partial factors for design according to mc 2020 and various ccs. cc γc γs γg self-weight γg other permanent loads γq (imposed) γq (wind) cc1 1.1 0.975 1.125-1.2 1.25-1.3 1.0 1.05 cc2 1.15 1.0 1.125-1.225 1.275-1.325 1.075 1.15 cc3 1.15 1.0 1.15-1.25 1.3-1.375 1.25 1.3 table 3. recommended values of partial factors for assessment of existing structures according to mc 2020. where a denotes a decisive geometrical property and the subscripts "k" and "d" refer to characteristic and design value, respectively. equation 7 assumes that: • the resistance is obtained as the linear product according to equation 1. • the characteristic value of the material property (commonly strength) corresponds to a 5% fractile of the respective lognormal distribution. • unity characteristic values are considered for resistance model uncertainty and conversion factor, respectively. • statistical uncertainty is reflected in the estimate of characteristic resistance. the same assumptions apply also for relationship (8) in section 4.2. table 2 provides the recommended values of the partial factors for design according to mc 2020 for low, medium and high consequence classes (cc1 to cc3 respectively); γg denotes the partial factor for permanent load effects and γq for variable load effects. a detailed discussion on the partial factors for load effects is beyond the scope of this contribution but selected values are provided in table 2 and in table 3 in section 4.2 to allow for comparisons with the design requirements of eurocodes in section 5. for the sake of brevity, the following discussion is focused on cc2 only. the partial factors for materials in table 2 are derived using relationship (7) and considering: • the probabilistic models provided in section 2. • a 50-year reference period along with the target reliability index β = 3.8 (cc2) and sensitivity factors αr = 0.8 for resistance and αe = −0.7 for load effects. the partial factors in table 2 are in broad agreement with the values provided by eurocodes. important for many concrete structures is that γg for self-weight can be reduced to 1.25. in contrast, the γq-value for wind action effects is increased. note that en 1990:2002 and pren 1990:2021 keep partial factors for materials fixed across ccs and differentiates by adjusting partial factors for unfavourable load effects only. the background documents for mc 2020 further demonstrate that similar values of the partial factors are obtained when the informative annual target reliability indices β given in mc 2020 are taken into account along with the sensitivity factors adjusted for a 1-year reference period αr = 0.7 and αe = 0.8. note that mc 2020 provides partial factors for a number of other design situations such as the cases with higher execution quality, effective depth considerably exceeding 200 mm and effective depth measured in the finished structure. the draft mc 2020 emphasises that the recommended values of the partial factors for design or assessment may only be applied when the conditions of the structure under consideration comply with the assumptions adopted therein. examples of the cases when the γ-values should be updated include: • changes of the target reliability level and/or of the sensitivity factors • decreased or increased variability of a basic variable such as a material property, geometry, load effect or model uncertainty • basic variables having other probabilistic distributions than those assumed when deriving the γ-values in particular for existing structures, the partial factors often need to be updated considering structurespecific information. higher-level methods should be applied when structure-specific conditions deviate from those commonly accepted. the more advanced approaches include the design (assessment) value method (using e.g. equation 7), a full-probabilistic approach (with updated models for basic variables and without making assumptions on the values of the sensitivity factors), or risk analysis (by which the target level can be optimised for the particular structure). 211 m. sýkora, w. botte, r. caspeele et al. acta polytechnica ctu proceedings 4.2. partial factors for assessment in order to identify the need for a safety measure to achieve a required reliability level at uls for the existing structure, a relationship similar to equation 7 is considered for the partial factor for concrete: γm = [exp(−1.645 vf is)]/[ µθr µac exp(−αr β √ v 2θr + v 2 ac + v 2 f is ] (8) where the in-situ concrete compressive strength evaluated from core tests, fc, is, is considered and the conversion factor is not included. table 3 provides the recommended values of the partial factors for assessment to verify the need of safety measure(s) (termed as "assessment of the existing structure" hereafter for brevity) under the assumptions provided below. the partial factors for materials in table 3 are derived using the equation 8 and considering: • the probabilistic models provided in section 2 along with vf c,is = 0.15 and vf y = 0.045, and various levels of the load effect model uncertainty (on which the γg-values are dependent while the γq-values are affected insignificantly) • an annual reference period along with the target reliability index β = 3.3 (cc2) and modified sensitivity factors αr = 0.7 and αe = −0.8 regarding decisive geometrical parameters, related uncertainties are described in mc 2020 for two cases: a) measurements in the finished structure and b) normal uncertainty. following section 2.2, the partial factors in table 3 are based on µa/anom = 1, vac = 0.015, and vd = 0.01 in case a). in case b), it is assumed that reliability verification adopts information on geometry from documentation that is checked by in-situ measurements, but the geometry of the decisive section is not directly verified. µac/ac,nom = 1, µd/dnom = 0.95, vac = 0.04, and vd = 0.05 are then considered. the draft mc 2020 specifies that the partial factors in table 3 should only be applied when: • concrete compressive strength is verified by adequate destructive tests (following standards on the evaluation of concrete in-situ strength such as en 13791:2019). • the type and amount of reinforcement is verified. • degradation does not affect structural behaviour significantly. additionally to the conditions described in section 4.1, updating of the recommended values may be needed e.g. due to structural damage including deterioration, differences in material properties, detailing provisions or execution tolerances. note that the partial factors for resistance may need to be increased to cover epistemic uncertainties related to the lack of knowledge about important details that may have a significant influence on structural behaviour such as actual geometry or amount of reinforcement in joints of frames. in these cases, a sensitivity analysis can help to explore the effect of such uncertainties on structural behaviour. as an example, assumptions for representative favourable and unfavourable situations might be made and structural analysis can then reveal the effect of this uncertainty on structural resistance. the draft mc 2020 further emphasises that human safety may require reliability levels higher than those provided in the mc for the assessment (such as annual β = 3.3 for cc2) if structural failure is expected to result in human losses. in particular, for special cases e.g. with many persons at risk or with large failure consequences, higher target levels might be more appropriate and the partial factors given in table 3 should be increased. it is seen from table 3 that all mc 2020 partial factors for the assessment are considerably lower than those provided in the mc and in eurocodes for design. as an example of the important change, γs reduces from 1.15 to 1.0. this decrease can be associated with two differences between design and assessment situations of similar importance: 1. the probability of the fractile associated with the design (assessment) value of resistance increases by an order of magnitude from φ(αr × β) = φ(0.8 × 3.8) = 1.18 ‰ for design to φ(0.7 × 3.3) = 1.04% for the assessment. 2. the effective depth is assumed to be measured in the existing structure and the bias and cov becomes more favourable than in design (cf. figure 2 indicating the difference of about 10% for normal uncertainty and measurements in the finished structure). in addition, mc 2020 again provides partial factors for other assessment situations (concrete strengths with low variability when vf c,is = 0.08 or reliability assessments based on information about geometry from verified documentation). 5. comparison of design and assessment requirements in this section, the requirements imposed by the recommended values of the partial factors according to eurocodes ("ec", primarily intended for design), draft mc 2020 for design (mc design), and for assessment (mc assessment) are compared. table 4 provides the overview of the values considered in this section, focusing on cc2 only. the comparison is made on the basis of a fundamental reliability condition, rd = ed, for compression of the short column with a dominating concrete contribution according to equation 9 and for bending of the beam with a dominating reinforcement contribution according to equation 10: 212 vol. 36/2022 probability of fib model code 2020 resistance variables reference γc γs γg permanent loads γq (imposed) γq (wind) ec* 1.5 1.15 1.35 1.5 1.5mc design 1.3** 1.85 mc assessment 1.15 1.0 1.25**,*** 1.075 1.15 *considered here to be primarily intended for design. **considering self-weight and other permanent actions equally important. ***averaged over low and normal levels of load effect model uncertainty, rounded. table 4. overview of partial factors considered in comparison of design and assessment requirements (cc2). ac fck/γc = γg gk + γq qk = γg(1 − κ) + γq κ = ed(κ) (9) as fyk dnom/γs = γg gk + γq qk = ed(κ) (10) where κ = qk/(gk + qk)is the load ratio based on the characteristic values of load effects; and ed = design (assessment) values of the load effect. the ratio of section or reinforcement area required by mc assessment to the requirement by eurocodes becomes: ρc,ex = ac,ex ac,ec = γc,ex ed,ex(κ) fck,is γc,ec ed,ec (κ) fck ≈ 1.15 γc,ex ed,ex(κ) γc,ec ed,ec (κ) (11) ρc,ex = ac,ex ac,ec = γs,ex ed,ex(κ) fyk dm γs,ec ed,ec (κ) fyk dnom ≈ 1.05 γs,ex ed,ex(κ) γs,ec ed,ec (κ) (12) in equation 11, ec is based on the characteristic value of concrete strength from control specimens, fck = µf c exp(1.645 vf c) ≈ 0.85 µf c considering vf c = 0.1. mc assessment applies the characteristic value of in-situ strength affected by the bias of the conversion factor. when the effect of concrete aging is-mostly conservativelyignored, the characteristic value might be estimated as fck,is ≈ µη µf c exp(1.645 vf c,is) which leads to fck,is ≈ 0.74 µf c for µη = 0.95 and vf c,is = 0.15, and to fck/fck,is ≈ 1.15. for mc design based on fck, this effect plays no role and equation 11 reduces to ρc,des = γc,des ed,des(κ)/[γc,ec ed,ec (κ)]. in equation 12, all three approaches are based on the same characteristic value of yield strength, fyk, and there is no numerical effect on ρs-values. ec is based on the nominal value of effective depth, dnom, while mc assessment assumes that in-situ measurements provide its mean, dm. according to section 2.3, it is then considered dnom/dm ≈ 1.05. for mc design based on dnom, equation 12 reduces to ρs,des = γs,des ed,des(κ)/[γs,ec ed,ec (κ)]. when evaluating ρ-values, the load ratio is a study parameter considered in the range from 0.1 to 0.7 characteristic for rc structures [14] and the values of partial factors are considered according to table 4. figure 3 displays the variation of ratio ρ for mc design. when ρ < 1, mc design leads to lower requirements than ec. in this case, there is no difference between flexural resistance and columns under compression as the same material factors are provided by ec and mc design (cf. table 4). in the case with imposed loads, mc design leads to nearly same levels as ec (the former being lower by 1-3%). for wind pressure (and similarly for snow loads), mc design (50 y.) requires higher resistances (5-15%) as the γw -value is significantly increased in comparison to ec. figure 3. variation of ratio ρ for mc design. figure 4 displays the variation of ratio ρ for mc assessment. in this case, mc assessment is below ec for all situations. in comparison to flexural resistance, 213 m. sýkora, w. botte, r. caspeele et al. acta polytechnica ctu proceedings figure 4. variation of ratio ρ for mc assessment. systematically lower levels are obtained for columns in compression. for flexural resistance, 75-80% of the reference ec level is mostly reached for variable actiondominated or permanent action-dominated structures, respectively. for columns, the level drops to 70-75%. for both failure modes, slightly lower levels are obtained for imposed loads. the ρ-ratios shown in figure 4 can be considered low and consequently the reliability levels associated with mc assessment might be questioned. such reduced requirements are attributed mainly to a lower target reliability adopted in the mc for assessment. as an example, for κ = 0.5 and wind pressure, the γc-value increases from 0.74 (obtained for β = 3.3) to 0.83 (β = 3.7). a smaller effect is related to reduced uncertainty as the measurements in the existing structure are assumed. 6. discussion besides the partial factors for design and assessment discussed in this contribution, the draft mc 2020 provides guidance for a number of other design and assessment situations. while new parts of an upgraded structure are to be designed in a way proposed for new structures, a specific approach for the reliability assessment of existing parts of the upgraded structure is provided. this approach is based on requirements adopted also for assessment in section 4.2, addressing the need of concrete core tests, the verification of the actual reinforcement, and a check of the degradation. however, in contrast to section 4.2, the target reliability levels for design are considered and dimensions are assumed to be measured in-situ in all cases (as is common practice before structural interventions). the draft mc 2020 provides detailed guidance for the specification of the target reliability levels considering annual or 50-year reference periods, somehow recommending the use of the former one. while both reference periods can be considered in design situations, in assessment situations mc 2020 focuses only on annual reference period and recommends annual target levels (requiring that the annual target reliability should be fulfilled in every year of a service life of the structure). for the two reference periods, different values of the sensitivity factors are recommended. 7. summary and concluding remarks the partial factors for materials, permanent load effects, and variable load effects are presented in draft fib mc 2020 in tables for the clusters of cases depending on consequence classes and variability of basic variables. formulas and background information facilitate updating of the recommended values, i.e. adjusting the partial factors for structure-specific conditions. the key inputs for deriving the partial factors include the probabilistic models of basic variables. different assumptions may then need to be adopted in design and assessment situations; the main differences are related to the considerations of the concrete strength measured in control specimens (design) and in-situ compressive concrete strength (assessment) as well as to the models for geometrical properties. the presented comparison of the requirements imposed by eurocodes and mc 2020 for design reveals on average insignificant differences (with somewhat increased partial factors for snow and wind loads and reduced partial factor for permanent actions in the model code). as a key new feature in the draft mc 2020, a lower target reliability is recommended for the assessment and the respective assessment requirements may decrease by about 25%. the conditions specified in mc 2020 should then be satisfied, including in-situ verification of concrete strength, reinforcement, and effects of degradation. an important principle when developing mc 2020 proposals was that the model code should be consistent with eurocodes, primarily with pren 1990 for basis of design and assessment and pren 19921-1 for design and assessment of concrete structures including bridges. regarding specific issues for concrete structures, mc 2020 often offers more detailed guidance than eurocodes. in particular, the revised mc 2020 will provide designers and code makers with wider possibilities to utilise actual data and long-term experience in assessments of existing structures. acknowledgements this study is based on the developments of fib com 3 tg 3.1. it has been partly supported by the czech science foundation under grant 20-01781s; results of grant ltt18003 supported by the ministry of education, youth and sports of the czech republic have been utilised. 214 vol. 36/2022 probability of fib model code 2020 resistance variables references [1] s. matthews, a. bigaj-van vliet, j. walraven, et al. fib model code 2020: towards a general code for both new and existing concrete structures. structural concrete 19(4):969-79, 2018. https://doi.org/10.1002/suco.201700198. 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[11] m. engen, m. a. n. hendriks, g. monti, et al. treatment of modelling uncertainty of nlfea in fib model code 2020. structural concrete 22(6):3202-12, 2021. https://doi.org/10.1002/suco.202100420. [12] cen/ tc250/ sc10/ ahg reliability (chair acwm vrouwenvelder). technical report for the reliability background of eurocodes (draft, june 2021), cen tc250/ sc10, 2021. [13] jcss. probabilistic assessment of existing structures (ed. d. diamantidis). joint committee on structural safety, rilem publications s.a.r.l., paris, 2001. [14] n. e. meinen and r. d. j. m. steenbergen. reliability levels obtained by eurocode partial factor design a discussion on current and future reliability levels. heron 63(3): 243-302, 2018. https://heronjournal.nl/63-3/2.pdf. 215 https://doi.org/10.1002/suco.201700198 https://doi.org/10.1002/suco.201900562 www.jcss.byg.dtu.dk https://doi.org/10.1002/suco.202000211 https://doi.org/10.1002/suco.201700244 https://doi.org/10.1002/suco.201900153 https://doi.org/10.1016/j.probengmech.2015.09.008 https://doi.org/10.1016/j.probengmech.2015.09.008 https://doi.org/10.1002/suco.201700169 https://doi.org/10.1002/suco.202100420 acta polytechnica ctu proceedings 36:206–215, 2022 1 introduction 2 probabilistic models for resistances 2.1 general 2.2 material properties 2.3 geometrical variables 2.4 model uncertainties 3 updating of probabilistic models 4 recommended values of partial factors in draft mc 2020 4.1 partial factors for design 4.2 partial factors for assessment 5 comparison of design and assessment requirements 6 discussion 7 summary and concluding remarks acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.36.0099 acta polytechnica ctu proceedings 36:99–108, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague artificial intelligence finite element method hybrids for efficient nonlinear analysis of concrete structures michael a. krausa, b, ∗, rafael bischofa, walter kaufmanna, b, karel thomaa a eth zürich, institute of structural engineering (ibk), chair for concrete structures and bridge design, stefano-franscini-platz 5, ch-8093 zürich, switzerland b eth zürich, center for augmented computational design in architecture, engineering and construction, design++ initiative and immersive design lab, stefano-franscini-platz 5, ch-8093 zürich, switzerland ∗ corresponding author: kraus@ibk.baug.ethz.ch abstract. realistic structural analyses and optimisations using the non-linear finite element method are possible today yet suffer from being very time-consuming, particularly in case of reinforced concrete plates and shells. hence such investigations are currently dismissed in the vast majority of cases in practice. the "artificial intelligence finite element hybrids" project addresses the current unsatisfactory situation with an approach that combines non-linear finite element models for reinforced concrete shells with scientific machine learning algorithms to create hybrid ai-fem models. the ai-based surrogate material model provides the material stiffness as well as the stress tensor for given concrete design parameters and the strain tensor. this paper reports on the current status of the project and findings of the calibration of the ai-based reinforced concrete material model. we successfully calibrated and evaluated k-nearest-neighbour, lgbm and resnet algorithms and report their predictive capabilities. finally, some light is shed on the future work of integrating the ai surrogate material models back into the finite element method in the course of the numerical analysis of reinforced concrete structures. keywords: concrete material model, machine and deep learning, nonlinear finite element method, surrogate modeling, uncertainty quantification. 1. introduction digital design and manufacturing methods, such as those to be developed at the new immersive design lab (idl) at eth zurich, offer great potential for significantly more efficient and sustainable construction. in order to ensure the structural safety, economic efficiency and sustainability of complex structures, reliable and powerful models for automatic analysis, optimisation and design are essential. however, such models are largely lacking to date. this is especially true for concrete structures, as their behaviour is highly non-linear. realistic structural analyses and optimisations using the non-linear finite element method (fem) are possible today, yet being very time-consuming even for the case of extraordinary computational capacities. for reasons of temporal and monetary efficiency, established traditional and in many cases excessively conservative design methods without structural optimisation are still used in the vast majority of projects in practice today. the increased public awareness of and demand for a sustainable built environment however urges civil engineers to make use of structural efficiency to the maximum level possible. to that end, a two-phased research program [1, 2] is proposed to address this unsatisfactory situation. in the first phase, non-linear fem for reinforced concrete plates and slabs developed at eth zurich [3–8] are combined with scientific machine learning (sciml) algorithms to create hybrid ai-fem models, which are expected to be much more efficient compared to established analysis methods both in terms of the computing power required and the reliability of predicting the load-bearing behaviour. in the second phase of this project, the ai-fem-hybrids will be used within a novel generative design process for accelerated yet realistic conceptual design of bridges [1, 2]. within a fe analysis, a material model has to provide on the one hand the material stiffness matrix and on the other hand the current stress state. in the novel hybrid artificial intelligence finite element (aifem) model implementation for material-nonlinear reinforced concrete slabs and plates elements, the mathematical description of the material model is performed using scientific machine and deep learning algorithms, which act as functional approximators of the stress-strain relationship and the stiffness-strain relationship based on numerical simulations. the basic data set for training, validation and testing of the ai algorithms is generated by extensive simulations of strain states for different reinforced concrete configurations with the material model for reinforced 99 https://doi.org/10.14311/app.2022.36.0099 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en m. a. kraus, r. bischof, w. kaufmann, k. thoma acta polytechnica ctu proceedings concrete as implemented in the mentioned usermat in ansys mechanical apdl on the basis of the cracked membrane model (cmm) in combination with a reissner-mindlin layer model. this paper describes the process of data generation and calibration of the ai-based reinforced concrete material model for different machine and deep learning algorithms for stress and stiffness tensor prediction. this paper is organised as follows: we first provide background on the materials and methods from computational mechanics and scientific machine learning used in this paper in sec. 2. sec. 3 reports and presents selected numerical results of the sciml-based substitute concrete material models w.r.t. their approximation accuracy and statistical qualities. sec. 4 presents a discussion of current findings and sec. 5 sheds light on the future steps of incorporating the sciml-based substitute concrete material models into the fem software ansys. 2. materials and methods this section reports on materials and methods for database generation, development of machine and deep learning (ml/dl) models, and the evaluation criteria for the performance of the developed models as a surrogate for a fem. 2.1. concept this research project explores the potential of using ml/dl surrogate models of mechanically consistent nonlinear finite element material models for reinforced concrete plates and shells, developed at eth zurich, as a data-driven yet physics-informed ai method within a fem workflow. the project is divided into two parts. the ai-fem-hybrids of phase one provide predictions of the material stiffness tensor as well as the stress tensor for a current strain state, where the sciml algorithms act as functional approximators of the stress-strain relationship and the stiffness-strain relationship, which was found to be applicable to a number of materials [9]. the basic data set for training, validation and testing of the ai algorithms is generated by an extensive simulation of strain-stressstiffness states for different reinforced concrete configurations with the material model for reinforced concrete as implemented in the mentioned usermat in ansys mechanical apdl on the basis of the cracked membrane model (cmm) in combination with a layer model based on reissner-mindlin plate kinematics. the training, validation and testing of the sciml algorithms for stress and stiffness tensor prediction are presented and discussed in the following, where special consideration is given to model uncertainty quantification to allow for a future eurocodecompliant use within the semi-probabilistic design philosophy. with the completion of the first phase, a validated software demonstrator of ai-fem-hybrids for the efficient non-linear analysis and design of reinforced concrete structures is available. in the second project phase, the ai-fem-hybrids approach will be extended to structural optimisation of concrete bridges. the pilot project in this phase focuses on concrete bridges with a geometry defined by a few parameters. while the case study is specific, the developed optimisation methodology is kept as general as possible to allow its application beyond parametric concrete bridge structures. at the end of the research project, a validated software demonstrator of the aife-hybrids will be available for the efficient non-linear analysis, design and optimisation of reinforced concrete structures. in the future, this will then allow for (i) realistic structural analyses taking into account non-linearities and (ii) structural optimisation to be carried out much more efficiently. the outcomes of this research are potentially of high interest to industry and engineering practice, as structural concrete is the most widely used construction material worldwide and incorporation of these ideas allows for more economic, yet sustainable and reliable design. to achieve the objectives for phase one of the research project, the following sequence of steps was conducted: (1.) define relevant feature x and target variables y for a reinforced concrete material model (cf. tab. 1) (2.) create the database {x; y} of fem simulations (3.) calibrate different ml and dl algorithms given the database (4.) assess different metrics for accuracy and predictive capabilities of the ml resp. dl models (5.) implement a selection of ml resp. dl models for implementation in ansys (6.) verify efficiency and accuracy at industry scale via computation of example problems from reinforced concrete design 2.2. fem material model data generation the dataset generation within the ai-fem-hybrid project was performed by nonlinear finite element analysis of a reinforced concrete structure utilising the cmm-usermat [3], a mechanically consistent material model for reinforced concrete implemented in ansys mechanical apdl, cf. figs. 1 and 2. combined with a layer element (shell181), the non-linear fem analysis of reinforced concrete structures as girders or shells are possible. as presented in [3–5], an excellent agreement between experimental results and fem analysis utilising the cmm-usermat is reported. in order to provide the necessary dataset for the training of the machine learning algorithms, a dummy girder bridge (cf. figs. 1 and 2) was arbitrary loaded incrementally until failure, which enables the export of 100 vol. 36/2022 ai-fem-hybrids for nonlinear concrete materials aa a a 45.00 m a a 5.00 m a a 2.25 m a a 250 mm a a 250 mm aa a a 400 mm a a 1'200 mm a a 495 mm a a 480 mm a a 437 mm q n q n a b c d b 0 l/2 l/2 q n y xz x z y 8'600[kn] [knm] v z m y [kn] [knm] n x t x a z b z l/2 l/2 c z d z (a) (b) x z y w (x = l/2, s 1 ) w (x = l/4, s 1 ) w (x = l/2, s 2 ) w (x = l/4, s 2 ) q n = 1'911 kn/m z y x q 1 = 200 kn/m 11'250 -11'250 -4'500 4'500 -50'625 aa a a = a a = a a = a a = a a = aa a a = a a = a a = a a = a a = z x yq1·b0/2 q 1 parameter: l b 0 h 0 t inf t sup parameter: t w t end a x a y a z figure 1. dummy girder bridge used within ansys to generate the database. features xi targets yi reinforcement layer (rela) normal stress in x-direction σx cmm usermat model (cmm) normal stress in y-direction σy reinforcement area as shear stress in xy-direction τxy reinforcement diameter ds stiffness component k11 effective reinforcement ratio for tcm model ρs,ef f stiffness component k12 reinforcement yield stress fy stiffness component k13 reinforcement ultimate stress fu stiffness component k21 reinforcement ultimate strain ϵsu stiffness component k22 reinforcement angle θs stiffness component k23 concrete compressive strength fcc stiffness component k31 concrete ultimate strain ϵcu stiffness component k32 normal strain in x-direction ϵx stiffness component k33 normal strain in y-direction ϵy shear strain in xy-direction ϵxy table 1. list of features and targets for data generation and calibration of ai-fem-hybrids. note that ϵxy is the tensorial component of the shear strain γxy , i.e., ϵxy =γxy /2. all necessary features (input parameters x: strain tensor, reinforcement properties, material constants, etc. acc. to tab. 1) and results (stress and stiffness tensors acc. to tab. 1) in every integration point (gausspoints) for all converged load steps. this approach enables the collection of a considerable number of data points within one non-linear fem computations, as in almost every integration point an individual data set exists. for simplicity, a dummy girder bridge was chosen to minimise the number of parameters, e.g. the diameter of the rebars or the material constants. as soon as the general framework of the ai-fem-hybrids is established (proof of concept), the results of any dummy experiment and the results of any fem analysis using the cmm-usermat can be used to train the machine learning algorithms. in total the data set consists of: • ns,wor = 10, 636, 800 samples of concrete elements without reinforcement (configuration "cf 0") • ns,wr = 443, 200 samples of concrete elements with reinforcement (configuration "cf 1") 2.3. data pre-processing the data generated in the previous section is used to develop the ml resp. dl models. the dataset consists of over 11 million data points with 14 feature and 12 target variables, cf. tab. 1. some dimensions of the feature dataset are categorical (reinforcement layer, cmm-usermat model), while the remaining features are all of numerical nature. for the categorical features we use one-hot-encoding. all dimensions of the target dataset are of numerical nature. none of the attributes has any missing data since the data is generated via fem analyses. the range of values 101 m. a. kraus, r. bischof, w. kaufmann, k. thoma acta polytechnica ctu proceedings x z y integration poinst shell element shell181 l l b 0 h 0 a y a z a x a x z x y y x z (a) (b) (c) figure 2. detail of the cmm finite element within ansys to generate the database. of the features as well as the targets are extremely different, hence we employ normalization of the data. for each numerical variable vi we deduce its mean m and divide it by its standard deviation s (yielding its z-score) using: zi = vi − µi σi (1) for sake of brevity of this paper, we omit reporting further statistical properties or histograms of the dataset, which would usually be delivered in the exploratory data analysis step. 2.4. machine learning model development in this study, three ml resp. dl algorithms for regression are investigated for predictive capabilities for stress and stiffness target tensors given the strain tensor together with concrete and reinforcement features: • k-nearest-neighbours (knn) • lightgbm • artificial neural networks: residual neural networks (resnet) while knn is a strictly data-driven approach and hence needs to store some of the simulation data instances, lightgbm as well as resnet are surrogate functions without the need to store the training data. due to reasons of brevity of this paper, only selected results and implementation details can be reported. the ml models are developed using sklearn, keras, and xgboost (extreme gradient boosting) python libraries and run on the eth euler cluster1. 2.4.1. knn the k-nearest-neighbours algorithm is a decision method originally developed for classification tasks [10]. at inference time, it compares the previously unseen samples to all instances in a database and assigns classes by means of majority voting of the k nearest neighbours, where k and the distance metric are hyperparameters. the algorithm was later extended to support regression tasks [11] by interpolating between the values of the k nearest neighbours (e.g. through inverse distance weighted average) to obtain values for new, unseen samples. 2.4.2. lightgbm lightgbm is a gradient boosting decision tree framework [12] that has found application in many different data mining, classification and regression tasks. the lightgbm algorithm variant used in this paper integrates a number of regression trees to approximate the dataset. it contains many of the advantages of other common gradient boosting decision tree algorithms, such as sparse optimization, parallel training, regularization, bagging, and early stopping. however, it grows trees leaf-wise by greedily choosing the leaf that will lead to the largest improvement. this tree-growth method, in combination with a histogram-based memory and computation optimization, make lightgbm considerably more computationally efficient than other frameworks. 2.4.3. artificial neural networks fully-connected feed-forward neural networks (ffnn) consist of one or more layers, where each node is connected to every node in the following layer [13]. despite being very straight forward architectures, ffnns are universal function approximators [14] and have the advantage of being easy to implement and efficient to train. on the other hand, the lack of implicit bias limits their expressiveness and capability to generalise. furthermore, they are prone to pathologies like vanishing gradients [15]. we therefore extend the ffnn by adding skip connections [15] as well as batch normalisation [16] and dropout [17] , cf. fig. 3. these additions allow us to build deeper architectures, which were shown to generalise better in practice [18]. we will hereafter be referring to this model by resnet. 2.4.4. hyperparameter tuning we employ bayesian optimization in order to avoid running extensive grid-search over the entire hyperparameter space [19][20]. at first, 25 random points 1https://scicomp.ethz.ch/wiki/euler 102 vol. 36/2022 ai-fem-hybrids for nonlinear concrete materials fu l ly c o n n e c t e d b a t c h n o r m a l is a t io n d r o po u t fu l ly c o n n e c t e d + figure 3. resnet: neural network with skip connections on the hyperparameter space are evaluated. gaussian processes then serve as prior distribution in order to approximate the unknown function and a posterior distribution is maintained as 75 more observations are made, where expected improvement (ei) is used as exploration strategy [21]. the hyperparameters together with optimization intervals are reported in tabs. 2, 3, and 4. hyperparameter range number of neighbours k [1, 10] power of minkowski metric p [1, 3] table 2. hyperparameters for training settings together with ranges as used for bayesian optimisation of knn hyperparameter range log-scaling learning rate [10−5, 10−1] yes max depth [3, 50] no min child weight [0, 10] no number estimators [30, 300] no number leaves [10, 100] no min child samples [10, 30] no table 3. hyperparameters for training settings together with ranges as used for bayesian optimisation of lightgbm hyperparameter range log-scaling learning rate [10−5, 10−1] yes depth [3, 50] no width [0, 10] no activation {relu, no leakyrelu} table 4. hyperparameters for training settings together with ranges for bayesian optimisation of resnets for the optimization procedure, we split the dataset into a training, validation and test set at ratios of (50; 25; 25)% of the whole dataset. the regressor is trained on the training set using the hyperparameters chosen by the bayesian optimization algorithm. once the model loss converged, it is evaluated on the validation data set and the result serves as feedback for the bayesian optimization step to refine its posterior and select new hyperparameters. finally, the test set is used to estimate the model’s capability of generalizing. 2.4.5. model quantitative performance metrics and ai surrogate model selection model selection refers to the process of seeking for a model in a set of candidate models, which delivers the best balance between model fit and complexity [22]. for regression modeling, the quantitative performance metrics used with this research are the mean absolute error (mae), root mean squared error (rmse), rsquared (r2) value as well as the data variance v ar (which is the square of the standard deviation (sd)), cf. tab. 5. the mae shows the average difference between the actual values of the output variable in the original data vs. the predicted output values via the ml models. the lower the mae, the more precise the performance of the model is in predicting future occurrences of the output. the rmse is defined as the standard deviation of the response variable. values of r2 range from 0 to 1, where 1 is a perfect fit, and 0 means there is no gain by using the model over using fixed background response rates. it estimates the proportion of the variation in the response around the mean that can be attributed to terms in the model rather than to random error. for comparing different models, we provide taylor diagrams [23] and additionally report r2, rmse and mae values. taylor diagrams are used to quantify the degree of correspondence between the modeled and observed data using the pearson correlation coefficient, rmse, and standard deviation. a model with the highest r2 and the lowest rmse and mae is preferred. the details of how mae, rmse, and r2 are calculated are shown in tab. 5. mae 1 n ∑n i=1 |ŷi − yi| rmse √ 1 n ∑n i=1(ŷi − yi)2 r2 1 − ∑ n i=1 (ŷi−yi)2∑ n i=1 (ŷi−ȳ)2 , ȳ = 1 n ∑n i=1 |yi| table 5. regression performance evaluation metrics. where n is the sample size, ŷi is the predicted and yi is the true target value, sse is the sum of the square of error and sst is the total sum of squares. 3. results this section describes the results obtained with knn, lightgbm, and resnet and evaluates their performance as a surrogate for the fem data generated and used within this study, cf. sec. 2.3. all three ml models were trained, validated and tested using the data from the total of 12 million 103 m. a. kraus, r. bischof, w. kaufmann, k. thoma acta polytechnica ctu proceedings data points. comparing the r2 values, all models have high values of around 0.98. this indicates that the ml algorithms are able to describe over 98% of variations in the data, and are highly predictive of the output based on the feature variables used in the study. regarding the mae and rmse, the knn model performs one order of magnitude better than the other two ml models for the cf 0 configuration (without reinforcement) at a level of 0.3% resp. 0.05%, while for the cf 1 configuration (with reinforcement) the resnets has the lowest rmse value of 2.4% and the lgbm has the lowest mae value of 1.8%. the results suggest that all ml models developed for predicting the stress (all in unit: mpa) and stiffness (all in unit: mnm2) tensor components are promising, with resnet being the most predictive model amongst the three. the model performances, described by the mean absolute error mae, root mean squared error rmse and r-squared have converged to a stable minimum during training. 3.1. knn model results fig. 4 provides selected results for the knn algorithm in the two configurations "cf 0" (without reinforcement) and "cf 1" (with reinforcement). it is important to note, that the size of the trained knn model is of around 1.13 gb, as it needs to store all training samples in order to make new predictions. 3.2. lgbm model results fig. 5 provides selected results for the lgbm algorithm in the two configurations "cf 0" (without reinforcement) and "cf 1" (with reinforcement). 3.3. resnet model results fig. 6 provides selected results for the resnet algorithm in the two configurations "cf 0" (without reinforcement) and "cf 1" (with reinforcement). 3.4. overall model comparison results the performances of different ai models according to the mentioned criteria are reported in tab. 6 for configuration "cf 0" (without reinforcement) and in tab. 7 for configuration "cf 1" (with reinforcement). model rmse mae r2 v ar knn 0.00045 0.00310 0.99955 0.00044 lgbm 0.00530 0.01837 0.99440 0.00560 resnet 0.00212 0.01576 0.99760 0.00238 table 6. performance of different models on test set without reinforcement for model selection purposes, we provide selected results for the taylor diagrams in the two configurations "cf 0" (without reinforcement) and "cf 1" (with reinforcement) in fig. 7. model rmse mae r2 v ar knn 0.06678 0.01579 0.93789 0.06678 lgbm 0.02783 0.02329 0.97390 0.02783 resnet 0.02417 0.03301 0.97715 0.02416 table 7. performance of different models on test set with reinforcement 4. discussion the choice for the three investigated ai algorithms was done under three aspects: modeling bias, computing efficiency and prediction precision. the knn and lgbm belong to the family of non-parametric models, whereas resnet is a parametric model. nonparametric ml algorithms promise greater performance at the cost of higher data requirements and training times together with a risk of overfitting. parametric ml/dl models on the other hand come with high modeling bias towards the functional relation, yet neural networks provide a great enough expressiveness for modeling the database. concerning computational efficiency in the prediction stage (which is called a lot of times during fem analysis in each iteration step) knn requires to store the dataset and call it at prediction time, leading to impractical lengthy procedures. lgbm and resnet however are much more time efficient and hence to be preferred for an implementation into the fem analysis. the ai models perform better in cf 0 (without reinforcement) than in cf 1 accross the different quality measures. looking at the r2 values, all models have high values of approx. 0.99 (cf 0) resp. 0.97 (cf 1). this means, that the ml algorithms are able to describe over 97% of variations in the data, and hence are highly predictive for the outputs. for cf 0 rmse is around 0.35% and mae is around 1.5%, while for cf 1 rmse as well as mae lie about 3% and are thus one magnitude bigger than for cf 0. it is noteworthy that on average, every target is predicted well, while the prediction’s standard deviation is dependent on the respective target quantity. given all model results, especially the stiffness tensor terms k22 and k23 show severe model predictive deviations. for model selection, inspection of the taylor diagrams in fig. 7 suggests that all described ai models possess great approximation quality with slight differences, where knn outperforms resnet and lgbm (in the order of decreasing approximation quality). the results in summary suggest that the ai models developed for predicting the stress and stiffness responses of the cmm user-mat are promising and may be used within a fem. given the outlines in the beginning of this section towards computational efficiency, the resnet and lgbm are chosen for further implementation into the fem analysis process. another interesting thought is to use the calibration of ai algorithms on such numerical datasets for non-intrusive verification purposes of the implementa104 vol. 36/2022 ai-fem-hybrids for nonlinear concrete materials (a) (b) (c) (d) figure 4. knn results for: (a) stress component σx in cf 0, (b) stiffness component k11 in cf 0, (c) stress component σx in cf 1, and (d) stiffness component k11 in cf 1. (a) (b) (c) (d) figure 5. lgbm results for: (a) stress component σx in cf 0, (b) stiffness component k11 in cf 0, (c) stress component σx in cf 1, and (d) stiffness component k11 in cf 1. 105 m. a. kraus, r. bischof, w. kaufmann, k. thoma acta polytechnica ctu proceedings (a) (b) (c) (d) figure 6. resnet results for: (a) stress component σx in cf 0, (b) stiffness component k11 in cf 0, (c) stress component σx in cf 1, and (d) stiffness component k11 in cf 1. (a) (b) (c) (d) figure 7. taylor diagrams for: (a) stress component σx in cf 0, (b) stiffness component k11 in cf 0, (c) stress component σx in cf 1, and (d) stiffness component k11 in cf 1. 106 vol. 36/2022 ai-fem-hybrids for nonlinear concrete materials tion of the ground truth data generating mechanism. especially the data points in far distance to the diagonal line in figs. 4, 5,6 give rise to outlier detection methods. it is currently under investigation whether the identified outliers are truly faulty data (i.e. the generating fem material model is not correct) or the ai algorithm with its current state is not able to approximate the respective quantities well. 5. conclusions and outlook this paper presented intermediate results of the first phase of a two-stage research project conducted currently at eth zürich. it reports on calibrating aibased surrogates for a highly nonlinear reinforced concrete plate and shell model upon the cmm. all three ai algorithms have proven to furnish as valid candidates for surrogate models to predict the stress and stiffness tensors used within a fem. the datadriven knn algorithm performed slightly better that the lgbm and resnet model for prediction standard deviation and rmse, whereas on average all models show almost full correlation. future research is concerned with further improving predictive capabilities by training an ensemble model using the knn, resnet and lgbm models. finally, the non-intrusive verification analysis of the ansys user-mat by ai surrogate modelling will shed light on generalisation of this idea of physics informed ml [24]. acknowledgements the authors would like to acknowledge the facilities at eth zürich as well as the design++ resp. immersive design lab for providing computational resources for the machine and deep learning parts of this research project. this research was funded through the eth foundation grant no. 2020-hs-388 (provided by kollbrunner/rodio). references [1] m. kraus, w. kaufmann, k. thoma. 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[24] r. bischof, m. kraus. multi-objective loss balancing for physics-informed deep learning. arxiv preprint arxiv:211009813 2021. 108 https://doi.org/10.1177/0049124104268644 acta polytechnica ctu proceedings 36:99–108, 2022 1 introduction 2 materials and methods 2.1 concept 2.2 fem material model data generation 2.3 data pre-processing 2.4 machine learning model development 2.4.1 knn 2.4.2 lightgbm 2.4.3 artificial neural networks 2.4.4 hyperparameter tuning 2.4.5 model quantitative performance metrics and ai surrogate model selection 3 results 3.1 knn model results 3.2 lgbm model results 3.3 resnet model results 3.4 overall model comparison results 4 discussion 5 conclusions and outlook acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0151 acta polytechnica ctu proceedings 38:151–161, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague a theoretical assessment of transport emissions from institutional buildings in a norwegian municipality oskar fahlstedt∗, rolf andré bohne norwegian university of science and technology, høgskoleringen 7a, 7034, trondheim ∗ corresponding author: oskar.fahlstedt@ntnu.no abstract. municipalities are key actors in reaching the sustainable development goals 9, 11, 12, 13 set by the united nations. the aim of this paper is to develop, on a theoretical level, a sustainable building refurbishment framework for municipalities. the focal point of the framework is reduction of embodied and operational carbon emissions deriving from buildings, including travel-induced emissions. the institutional buildings governed by a norwegian municipality are the case study objects in this study. the framework that is currently is used in the municipality value the building stock based on the quality and cost of refurbishing them. no framework currently includes travel-induced emissions in the evaluation. the intention is to provide a holistic framework that includes this aspect when deciding if a building should be refurbished or demolished. the omission of transport emissions can lead to truncation errors in assessments. the transport to and from the school generates higher emissions in certain scenarios. based on the knowledge in the assessment, a theoretical framework that assesses this is developed. keywords: co2 abatement, travel-induced emissions, location, institutional buildings, municipality. 1. introduction 1.1. energy refurbishments the built environment is responsible for as much as 70 % of human-induced greenhouse gas emissions (ghg) [1, 2], and the united nations (un) predicts that 70 % of the world population will live in cities by 2050 [1, 3]. additionally, predictions are that 85– 95 % of european buildings in operation today will still be in operation by 2050, and most are of inadequate energy standards [4]. thus, they are responsible for a large share of total carbon emissions and sensible actions are refurbishing existing building stocks to lower energy use and indirect carbon emissions. there is no lack of research providing deep-refurbishment methods and frameworks [4, 5]. however, yearly refurbishment rates are at 1 % across europe, and for deep refurbishments (refurbishment intended to lower energy use by a minimum 60 %), rates are between 0–0,2 % [4]. furthermore, an overwhelming body of literature focuses on optimizing on a building scale. thus only emphasizing mitigation of embodied and operational emissions whilst there are reasons to believe that other emission sources are significant when the scope expands from building to urban scale [6]. 1.2. transport emissions and buildings anderson et al. [5] calculate the emissions for three different building types in munich, germany. the authors adopt a process-based (streamlined) lca framework and “expand” it to capture a new impact category they describe as “induced impacts”, which refer to impacts deriving from embodied and use phase emissions from transportation. the buildings are typical for each location in the city which are multi-family houses (city center), row houses (city periphery), and single-family houses (districts / rural areas). the study analyzes the embodied and operational emissions deriving from buildings and user transportation. the scope is not limited to operational transport emissions, as the embodied emissions for transport and infrastructure can represent a large share of total transport emissions. the calculated induced impacts represent 46 to 54 % of total emissions. furthermore, after a sensitivity analysis, it is concluded that moving any of the three building types further away from the city center will increase total emissions [5]. yu et al. [7] argue that building and city scale has been highly researched, while studies on urban precincts (e.g., districts) are not sufficiently covered in the literature. the authors state that assessment on this scale should include [7, p. 2] “1) embodied emissions that are associated with the construction, maintenance and end-of-life treatment of all precinct objects; 2) operational emissions that are generated from the operation of all precinct objects, 3) occupantrelated transport emissions that are generated from daily commuting, business trips and personal travels”. a precinct in adelaide, australia, is the case study, and occupant-related transport and personal travels have the largest share of emissions (40 %), followed by building operational (32 %) and embodied emissions (28 %). anderson et al. [5] and yu et al. [7] argue that it is important to include mobility emissions. fenner et al. [8] share the same opinion and perform a life-cycle 151 https://doi.org/10.14311/app.2022.38.0151 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en oskar fahlstedt, rolf andré bohne acta polytechnica ctu proceedings carbon emission assessment (lcco2a) to calculate embodied, operational, and commuting emissions for a campus building in the usa. however, without considering the embodied impacts of transport as they are perceived as negligible compared to operational emissions. the authors conclude that the emissions deriving from operating the building are the highest. although, not without reservations for a different outcome if a building with the latest energy standards is assessed instead [8]. in a study performed in norway, lausselet et al. [9] evaluated the application of zero-emission neighborhoods (zen). they identify the operational phase and mobility of inhabitants as the most important factors for transforming existing neighborhoods into zen but also highlight other important factors (see [9]). bastos et al. [10] investigate the importance of including mobility when assessing buildings in lisbon and discovered that emissions from transportation can exceed the emissions deriving from the building itself. the study investigates one apartment building in the city center and one semi-detached house in a suburban area. the suburban house has the highest emissions, and relocating it to the city center would reduce emissions significantly. thus, in the investigated scenario, it makes more sense to focus on transportation rather than building emissions [10]. in another study by drouilles et al. [11] is the energy efficiency of five buildings in switzerland’s urban, peripheral, and rural areas is analyzed against national energy targets. the results reveal that energy use in the buildings is twice as high as the national goals. furthermore, including daily mobility of occupants to and from buildings increase emissions by 40 % in the investigated areas. 1.3. the goal of the study no silver bullet exists for reducing emissions in the built environment. instead, it is necessary to implement a wide array of solutions holistically with urban and building research working in synergy to lower carbon emissions from the built environment. therefore, the work of governing bodies such as municipalities is crucial. they withhold a significant mitigation potential that is necessary to utilize for reaching the climate goals. the european union (eu) is striving to be carbon neutral in 2050, and norway is aligning with the goal as well [12–14]. the energy mix is a determinant factor for reaching this goal, and every country adheres to its own set of prerequisites for producing renewable energy [15]. thus, proven methods are not necessarily generic, and instruments to reach carbon neutrality are most likely country or even region specific. the literature has demonstrated the importance of transport emissions in holistic assessments of buildings [5, 7–11, 16]. the objective of this paper is twofold. first, investigate the importance of induced travel emission in a refurbishment strategy with a norwegian municipality as a case study. second, develop a theoretical framework as a suggestion for how to travel-induced emissions could be included together with operational and embodied emissions when finding refurbishment strategies for existing building stock. to guide the research, two questions are asked: (1.) how important is it to include transport emissions in refurbishment and new building assessments in countries that rely on low-carbon electricity? (2.) how can transport emissions be included in a refurbishment framework? 2. materials and methods 2.1. case study – bærum municipality a case study approach was used in a norwegian municipality when documents, building audits, strategies, and cost calculations were scrutinized. important information was also obtained through conversations with municipal officials and consultants managing the building stock. 2.1.1. background bærum municipality is located in the periphery of the norwegian capital oslo. hence, it is an attractive area to live in because of the work opportunities offered and the short commuting distance to the capital. thus, it has seen a steady population growth [17]. the population in the municipality is close to 130 000 [18]. the urban morphology is suburban and rural with detached houses although, some areas are densely populated with multifamily homes (mfh) and commercial buildings [17]. in regards to housing, the composition is 29 % single-family homes (sfm), 13 % semi -detached homes, 17 % row-houses, 3 % other, and 37 % mfh. there are over 100 000 light-duty vehicles registered in the municipality, with primary fuel sources of the cars being 28 % gasoline, 25 % diesel, 23 % electric, and 24 % other (mainly hybrid cars) [18]. as of 2016, 63 % of total emissions derive from light-duty vehicles (ldv), and with the inclusion of heavy-duty vehicles (hdv), it increases to 87 %. this is significant compared to the 3 % from heating the buildings, 4 % from construction machinery, and 6 % other (agriculture, waste management, energy supply, and anaerobic digestion from landfills) that constitute the remaining 13 % of total emissions [19]. the municipal real estate portfolio comprises a vast range of buildings with a combined area of 800 000 m2 to an insured value of 17.6 billion nok excluding the value of infrastructure and land [20]. nevertheless, the population is predicted to grow 11 % by 2030 and increase demand for public buildings. this is a concern for the municipality due to the scarcity of publicly owned land to build on [17, 19–21]. in 2014, the municipality appointed an external consultancy firm to evaluate the condition of the existing building stock, and they concluded that, on average, 24 % of the buildings achieve satisfactory 152 vol. 38/2022 a theoretical assessment of transport emissions . . . figure 1. abc-framework adapted from svein bjørberg in [20]. standards [20, 22]. the buildings were evaluated according to the norwegian standard for “condition survey of construction works – content and execution” (ns3424:2012), and results are reported on a consequence scale from 0–3. a score of 1.0 or less states that there are only minor consequences to be expected considering the standard of the building, but if the receives a score between 2.0–3.0, then there is a risk of medium to severe consequences. they are, for example, related to safety (structural safety, fire safety), health/environment (indoor air quality), aesthetics (surfaces), and economy (renovation, maintenance) [23]. according to the results from the study, average building condition is lower than comparable municipalities, and there is also a discrepancy within the range of buildings not reaching acceptable standards [20, 22]. the municipality is currently categorizing buildings into three levels (a, b, or c) depending on the total score from a quality assessment tool [20]. they are analyzed based on technical quality, adaptability, functionality, plot, and surrounding area. thus, it provides the municipality with an overview of the building stock that makes known where actions are needed to retain and increase the value of their buildings. furthermore, buildings categorized as a-level buildings in the framework are perceived to be worth the investment to reach the adequate standard demanded to align with the sustainable building strategy. the investment to upgrade from c to a is too high. thus, c-level buildings are kept until they pass their functional lifetime. after that, they are liquidated or demolished, and the land can fulfill other purposes [20, 22]. buildings categorized to b-level are under evaluation if they should be upgraded into a-level or downgraded to c-level instead. thus, they are important since categorizing a building into the wrong category can result in a non-sustainable solution [20]. school gfa [m2] hfa [m2] oe [kwh] 1 4639 3935 4,98e+05 2 7662 6170 9,66e+05 3 7215 6390 1,13e+06 4 5933 5084 7,60e+05 table 1. information about the four b-level schools. 2.1.2. institutional b-level buildings four schools previously categorized by the municipality as b-level buildings were assessed. the schools are of varying sizes, with indoor climates deemed inadequate, and none of the buildings have been refurbished in the last 25 years. none of the schools falls below 1.6 in the quality assessment [20, 23]. schools 1, 2, 3, and 4 house 298, 529, 117, and 525 children, respectively. the approximate cost for refurbishing them is 81 million nok and school 1 is the most expensive to refurbish. the children in the municipality are assigned to the schools based on the school district their homes are located in. school 3 is for children with special needs. thus, it does not belong to a specific school district since children are assigned based on their individual needs. information regarding the gross floor area (gfa), heated floor area (hfa), and operational energy (oe) use for 2019 are provided in table 1. 2.2. adapting the abc-framework figure 1 presents the adapted framework, with each line representing a requirement for the buildings. the red line in figure 1 illustrates building degradation over time to different levels followed by refurbishment actions. moreover, ideally, the school building should reach a higher quality level afterward. all school buildings that fall below c-level should be liquidated or demolished at some point as no sustainable investment is viable, and a downtrend arrow 153 oskar fahlstedt, rolf andré bohne acta polytechnica ctu proceedings figure 2. example of assumed distances between homes and schools at 1.5 km radius. illustrates the deterioration. the minimum requirement of the building is to achieve the norwegian technical standard (tek17) [24], and an a-level building should ideally surpass that minimum requirement. the vector for b-level buildings slightly increases in quality over time due to technical requirements becoming more stringent. the dotted line illustrates the minimum requirement from the norwegian technical standard for school buildings. the green dotted line represents the state-of-the-art school building with the best energy efficiency and lowest emissions. thus, as time progresses, the demand for higher functional quality increases since knowledge and technology advance. ideally, the municipality refurbishes a b-level school to a-level that performs better than the requirements in the technical standard. 2.2.1. building emissions the energy use from 2019 for each school was available in the case study. however, no information regarding the school’s embodied emissions was available. thus, allocating the share of emissions from the existing school buildings is difficult. therefore, the impacts from the previous production and construction stages were disregarded in this study. an archetype was used instead. it was based on average data from schools 1, 2, and 4, resulting in a gross floor area of 6078 m2 and a heated floor area of 5063 m2. it is assumed to be a b-level where 524 children attend the school. two improvement scenarios were developed for the school and are presented below: (1.) scenario 1: building a new school building that performs 20 percent better (88 kwh/heated m2a) than the minimum energy requirement for schools in [24]. embodied emissions (a1–b5) are assumed to be 6.9 kg co2eq/m2a. it is assumed that the school demand refurbishment after 30 years and embodied emissions are 65 percent lower (2.42 kg co2eq/m2a) at that point [25, 26]. even more stringent energy requirements are assumed to be in place after 30-years making the school 40 percent more efficient (66 kwh/heated m2a) than the requirement in the exiting energy standard for schools [24]. the end-of-life phase was not included in the assessment. (2.) scenario 2: a refurbishment scenario with the same energy requirements (88 kwh/heated m2a) as in the new building scenario. the embodied emissions are assumed to be lower (2.42 kg co2eq/m2a) since less material will be needed when refurbishing than constructing a new building. the refurbished building demands an additional refurbishment after 30 years with additional embodied emissions (2.42 kg co2eq/m2a). the school must adhere to the more stringent energy requirements (66 kwh/heated m2a) as the scenario with the new building. the electricity mixes used in the study are from the norwegian energy market. the first electricity mix is based on delivered electricity from the norwegian grid without any guarantees of origin (gos), and it has the highest carbon content in this study [27]. the second is a prediction for the average carbon content in norwegian electricity between 2015–2075 when importing from the european continent. third, the same scenario as the previous one, although electricity is only produced in norway [28]. the fourth is a low carbon scenario based on the reported average carbon content in the norwegian electricity mix during 2020 [29]. (1.) 402 g co2eq/kwh [27] (2.) 136 g co2eq/kwh [28] (3.) 18 g co2eq/kwh [28] (4.) 8 g co2eq/kwh [29] 154 vol. 38/2022 a theoretical assessment of transport emissions . . . 2.2.2. transport emissions a highly theoretical approach was used for the transport scenarios when it was assumed that all households were within a 2.5 km radius of the school. no specific routes were analyzed, and for simplicity, trip lengths were divided into 0.5, 1.0, 1.5, 2.0, and 2.5 km. as demonstrated in figure 2, schools 3 and 4 from the case study are located in such a manner that when the uptake radius is 1 km and increasing, they intersect one another. thus this is not a correct approach, but for simplicity, this approach is used as a theoretical exercise. it was assumed that round trips were made, meaning the children were dropped off and picked up each school day. the transport modes were limited to the car (gasoline, diesel, and electric) and bus (diesel and gas). the 2019 values for carbon emissions [g co2eq/pkm] deriving from transport using fossil fuels in norway were extracted from [30]. no specific adjustments to the transportation modes regarding modal share or improved fuel efficiency were made. the electric vehicles were assumed to use the same electricity mix as the buildings. for each distance, a share of students going by car or bus was assumed. these were between 10–50 % for each transport distance. 2.2.3. calculations the calculations was performed as follows: t ben = (em1 × gf a) + (oer1 × elm × hf a) + (em2 × gf a) + (oer1 × elm × hf a) (1) t ber = (em2 × gf a) + (oer2 × elm × hf a) + (em2 × gf a) + (oer1 × elm × hf a) (2) t t e = s × sd × t u × dd × 2 × et m. (3) abbreviations: em1 embodied emissions new building [kg co2eq/m2a] em2 embodied emissions refurbishment [kg co2eq/m2a] gf a gross floor area [m2] oer operational energy requirements [kwh/m2a] elm electricity mix [kg co2eq/kwh] hf a heated floor area [m2] s number of students [p] sd school days a year t u transport utilization [%] dd driving distance [km] et m emissions transport mode [kg co2eq/pkm] the t be represents “total building emissions”, and the “n” denotes the new building while “r” is the refurbishment scenario. these are the total embodied and operational emissions deriving from the building. for the new building, embodied and operational emissions are divided into two periods (30 years each) since the embodied emissions are after construction and lower after refurbishment. in the refurbishment scenario, the lower embodied emissions are calculated for both periods, while the operational phase differs. this is because the oer is assumed to be stricter once the building needs to be refurbished again after 30 years. t t e stands for “total transport emissions” and represents total emissions deriving from the operational phases of the selected transport mode. as mentioned earlier, distances to the school were assumed to be in the range of 0.5–2.5 km and utilization of the transport mode in the range 10–50 %. furthermore, the operational transport emissions were subtracted from the building emissions. to get an easy overview, tables were made (see table 2, 3 and 4) and a negative value marked in red means that operational emissions from transport are higher than embodied and operational from the buildings. 3. results the overall results showed that building emissions are always higher than the ones from transport for the first two electricity mixes (i.e., 402 and 136 g co2eq/kwh). the results changes once the building is supplied with low-carbon electricity mixes (18 or 8 g co2eq/kwh). in figure 3 are total emissions from the school building provided. the results reveal the significance of embodied emissions when the school is supplied with a low-carbon electricity mix. as seen, the new school generates higher emissions over its lifetime due to the initial embodied emissions. the gap between total emissions increases in the scenarios where low-carbon electricity is used compared to the two higher alternatives. distance users 10 % 20 % 30 % 40 % 50 % gasoline car [kg co2eq] in 8 g co2eq/kwh scenario. 2.5 km 4.27e+05 3.06e+05 1.86e+05 6.52e+04 -5.53e+04 2.0 km 4.51e+05 3.54e+05 2.58e+05 1.62e+05 6.52e+04 1.5 km 4.75e+05 4.03e+05 3.30e+05 2.58e+05 1.86e+05 1.0 km 4.99e+05 4.51e+05 4.03e+05 3.54e+05 3.06e+05 0.5 km 5.23e+05 4.99e+05 4.75e+05 4.51e+05 4.27e+05 table 2. in the table above, the results for the refurbishment scenario for the first 30 years presented. the emissions from transportation to and from the school are only higher (see red marking) than the embodied and operational emissions from the building if half of the children go by car over a distance of 2.5 km or greater. 155 oskar fahlstedt, rolf andré bohne acta polytechnica ctu proceedings figure 3. total emissions from the school building for the two scenarios with each of the four electricity mixes analyzed. distance users 10 % 20 % 30 % 40 % 50 % gasoline car [kg co2eq] in 8 g co2eq/kwh scenario. 2.5 km 4.00e+05 2.80e+05 1.59e+05 3.85e+04 -8.20e+04 2.0 km 4.24e+05 3.28e+05 2.31e+05 1.35e+05 3.85e+04 1.5 km 4.48e+05 3.76e+05 3.04e+05 2.31e+05 1.59e+05 1.0 km 4.72e+05 4.24e+05 3.76e+05 3.28e+05 2.80e+05 0.5 km 4.96e+05 4.72e+05 4.48e+05 4.24e+05 4.00e+05 diesel car [kg co2eq] in 8 g co2eq/kwh scenario. 2.5 km 4.13e+05 3.06e+05 1.99e+05 9.17e+04 -1.56e+04 2.0 km 4.35e+05 3.49e+05 2.63e+05 1.77e+05 9.17e+04 1.5 km 4.56e+05 3.92e+05 3.28e+05 2.63e+05 1.99e+05 1.0 km 4.78e+05 4.35e+05 3.92e+05 3.49e+05 3.06e+05 0.5 km 4.99e+05 4.78e+05 4.56e+05 4.35e+05 4.13e+05 table 3. in the table, the results are based on the refurbishment assumptions after 30 years presented. 3.1. first 30 years-low carbon mix the results for the first 30 years are presented in table 2. the values in the tables present the total emissions [kg co2eq] when t t e is subtracted from each t be scenario. the operational emission from transportation is only higher than the emission from the school if at least 50 % of the children travel by car while having a one-way distance of 2.5 km. furthermore, it only applies to the refurbishment scenario since emissions from the new building are always higher. 3.2. remaining 30 years-low carbon mix the embodied and operational emissions are the same for the remaining 30 years for both the building scenarios (see table 3). the emissions from transportation by gasoline cars are the same as in table 2 although higher since more time has passed. furthermore, since the embodied and operational emissions are assumed to be even lower, the emissions from the diesel car exceed the building emission given that 50 % travel 5 km (2.5 km to and from the school) each day. the results for the whole 60-year period are in table 4, and it shows that emissions deriving from the car only exceed emissions from the school if at least distance users 10 % 20 % 30 % 40 % 50 % gasoline car [kg co2eq] in 8 g co2eq/kwh scenario. 2.5 km 8.27e+05 5.86e+05 3.45e+05 3.45e+05 -1.38e+05 2.0 km 8.75e+05 6.82e+05 4.89e+05 4.89e+05 1.04e+05 1.5 km 9.23e+05 7.79e+05 6.34e+05 6.34e+05 3.45e+05 1.0 km 9.71e+05 8.75e+05 7.79e+05 7.79e+05 5.86e+05 0.5 km 1.02e+06 9.71e+05 9.23e+05 9.23e+05 8.27e+05 diesel car [kg co2eq] in 8 g co2eq/kwh scenario. 2.5 km 8.53e+05 6.39e+05 4.24e+05 2.10e+05 -4.41e+03 2.0 km 8.96e+05 7.25e+05 5.53e+05 3.82e+05 2.10e+05 1.5 km 9.39e+05 8.11e+05 6.82e+05 5.53e+05 4.24e+05 1.0 km 9.82e+05 8.96e+05 8.11e+05 7.25e+05 6.39e+05 0.5 km 1.02e+06 9.82e+05 9.39e+05 8.96e+05 8.53e+05 table 4. the table above presents the results for the whole (60 years) period. 50 % go by car over a distance greater than 5.0 km (2.5 km to and back from the school). 3.3. sensitivity analysis the results revealed that transport emissions only exceed building emissions under specific circumstances. the impacts from the operational phase of the school are already low, making the embodied emission decisive. thus, lower embodied emissions were assumed instead. in the original scenario, embodied emissions from refurbishments are 65 percent lower than the reference value of 6,9 kg co2eq/m2a obtained from literature [26]. the case study revealed that the energy use of school 1 is just above the requirements in tek17 [24]. only minor refurbishments were assumed instead, and embodied emissions were set to be 80 percent lower than the reference value. a low-carbon electricity mix (8 g co2eq/kwh) scenario was tested with gasoline and diesel cars. the results for the whole period (i.e., 60 years) are presented in figure 4. the figure shows the results for the first 30 years (left side), after another 30 years (middle), and the total emission on the right-hand side. the share of car users in the diagram (30–50 percent) with distances between 1.5–2.5 km is when transport emissions can be 156 vol. 38/2022 a theoretical assessment of transport emissions . . . figure 4. in the figure are the result for the first 30 years (left), the remaining 30 years (middle), and total emissions (right-hand side) when embodied emissions are reduced 88 % compared to the reference value for the first period presented. higher. since the role of embodied emissions was important, another analysis was performed. now minor refurbishment (80 percent lower embodied emissions) was assumed for the first period. however, for the second period, it is refurbished to align with the energy requirements (66 kwh/heated m2a) more material is needed. thus, embodied emissions were assumed to be 4.36 kg co2eq/m2a which follows guidelines for the use of “climate friendly” materials [25]. the results (see figure 5) reveal that car use emissions exceed the building emissions under specific circumstances during the first 30 years. once the building is refurbished after 30 years, the transport emissions never exceed the ones from the building. even if the electricity mix with the lowest carbon content is supplied to the building. 4. discussion 4.1. expanding the scope from building to urban scale the situation in the assessed municipality is similar to the rest of europe. it manages a large building stock, and refurbishments are needed [4]. with the scarcity of land and a growing population in the municipality, the argument for refurbishment is even stronger [19]. the literature reported that emissions from user transport should be included in assessments since they make a significant share of the total emissions [5, 7, 9, 10]. however, emissions deriving from user transport to and from the schools were only higher when the embodied and operational emissions were assumed to be very low, and half of the children are driven to school each day over a distance greater than 2.5 km. thus, the transport was less significant than predicted when considering previous studies [5, 7, 9, 10]. fenner et al. [8] reported that operational emissions were higher than transport emissions. however, the authors highlighted that their results could be different with a high-performing building scenario. this study investigated a high-performance building scenario, and results showed that building emissions were higher except under certain conditions. the school had to be supplied with a low-carbon electricity mix while having low embodied emissions. the aspect of embodied transport emissions was excluded in this assessment as well as in the one by fenner et al. [8]. however, anderson et al. [5] included them and showed that they are significant and thus ought to be considered in future assessments. the theoretical approach taken in this study means that it has limitations. for example, the archetype school was based on average values, while transport routes and modes were based on assumptions. future studies should aim towards finding applicable refurbishment packages and new building scenarios for one of the existing b-level schools. the case study discovered that schools were divided into districts so that transport routes could be analyzed in greater detail. thus, indicating the benefit of assessing schools. this is because households are confined within a specific area which means that it is possible to calculate accurate transport distances to their assigned school. this shows the strength of this approach since it enables holistic perspectives with solutions that help to reduce the total emissions on a municipal scale. combining the framework the municipality in the case study currently uses with the approach taken in this study avoids sub-optimization. the abc-framework identifies and sets the goals for the exiting stock whilst expanding the scope to include transport avoids carbon lock-ins. moreover, using the school districts makes it possible to optimize the location of a school within a district and evaluate if it is better to refurbish or build a new building at another location. 157 oskar fahlstedt, rolf andré bohne acta polytechnica ctu proceedings figure 5. in the figure are the result for the first 30 years (left), the remaining 30 years (middle), and total emissions (right-hand side) when embodied emissions are assumed to increase due to more extensive refurbishments needed after 30 years. 4.2. potential barriers for implementation the assessment in this study aimed to develop a theoretical framework, and more work is required. higher accuracy is achieved using bottom-up methods with on-site measurements [8, 31]. nevertheless, bærum municipality manages a large stock [20], and considering the urgency to reduce carbon emissions [12–14], it is not an economically sustainable strategy to address every building individually. therefore, it is necessary to decide a break-even point where the strategy should move towards using the top-down method based on statistical data to create archetypes [5] or potentially achieve more accurate results through clustering techniques [32]. murray et al. [32] claimed high accuracy of using clustering techniques to find archetypes of the swiss building stock. however, the techniques demand knowledge, and if it is not accessible among municipality officials, external support is needed, which can be costly. this potential barrier of how to transfer knowledge from the scientific community to decision-makers (dms) and city planners was not covered in the literature. thus, it should be investigated in future studies. there is some ambiguity about whether the current abc framework demands deep user knowledge. if the dms consider it to go beyond their skill level, a sophisticated black-box model is potentially beneficial as it theoretically can achieve high accuracy without demanding extensive technical knowledge about building physics. the municipality could then decide upon kpis, with data about building characteristics and location generating potential emissions based on the input. another option is to extrapolate data from a small sample or to use statistical data to create archetypes for their stock and then run energy simulations to find suitable refurbishment packages, which are methods already present in literature [11, 31]. however, complex machine learning techniques are often used to manage the computational power, which in turn demands extensive theoretical knowledge of the users [16]. 5. conclusion in section 1 two research questions were asked, and this section intends to answer them. the built environment is heterogeneous and varies depending on region, municipality, city, neighborhood, and building type, which is also the situation in bærum municipality. therefore, implementing holistic life-cycle frameworks covering both building and urban scales is recommended. it will aid dm in finding emission abatement strategies with less risk for carbon lock-in due to sub-optimization at the building scale. the results from this study reveal that the induced emissions from transportation can exceed embodied and operational building emissions. however, this was a theoretical exercise with a fictional school building, and further research with better data is needed. a sensitivity analysis demonstrated the importance of embodied emissions once a building uses low-carbon electricity. moreover, in the assessment, the hypothesis that transport emissions would be higher than building emissions was not the case. however, the work indicates the potential for higher reductions of ghg-emission on an urban scale when the method is applied. finally, the social aspect was not considered, but it should be in the future. for example, the economic situation determines the modal choice, and what the user perceives as valuable will set other requirements that ultimately influence emissions. the literature findings and information obtained from the case study have led to the development of a theoretical framework (see figure 6). it consists of 158 vol. 38/2022 a theoretical assessment of transport emissions . . . figure 6. theoretical framework based for holistic assessment of refurbishment strategies for existing buildings. three main emission drivers, the location of the building, the building itself (embodied and operational energy), and purchasing power (i.e., “social” for now) of inhabitants within each of the catchment areas of the schools. it is a suggestion for future detailed assessments. the social aspect was added impromptu due to the idea that household income will be deterministic for the accessibility of certain modes of transport. however, this aspect needs further development before being implemented hence the dotted line. the aforementioned emission drivers then determine the necessary inventory data for the calculation of environmental impacts. furthermore, emissions deriving from transportation are measured against emissions from refurbishing the building. a sensitivity analysis should be performed in situations when transport emissions are higher than the ones from the investigated building(s). this since greater reduction potential might be achieved if the services of the building (institutional in this case) are established at another location. for the school buildings, it insinuates that the school should be moved to a more suitable location. to move a building is drastic. it foremost refers to moving the service provided to location optimization strategies to reduce emissions that are not limited to the building scale but include transport from building users. references [1] un-habitat. the strategic plan 2020-2023, 2019. 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[30] statistisk sentralbyrå. co2-utslipp per person-/passasjerkilometer fordelt på kjøretøy og drivstoff, g co2/pkm. 2010-2019, 2021. [2022-02-22], https://www.ssb.no/transport-og-reiseliv/ artikler-og-publikasjoner/mindre-utslipp-fraveitrafikk-fly-og-tog?tabell=439538. [31] m. hu. life-cycle environmental assessment of energy-retrofit strategies on a campus scale. building research & information 48(6):659–680, 2020. https://doi.org/10.1080/09613218.2019.1691486. [32] p. murray, j. marquant, m. niffeler, et al. optimal transformation strategies for buildings, neighbourhoods and districts to reach co2 emission reduction targets. energy and buildings 207:109569, 2020. https://doi.org/10.1016/j.enbuild.2019.109569. 161 https://www.ssb.no/transport-og-reiseliv/artikler-og-publikasjoner/mindre-utslipp-fra-veitrafikk-fly-og-tog?tabell=439538 https://www.ssb.no/transport-og-reiseliv/artikler-og-publikasjoner/mindre-utslipp-fra-veitrafikk-fly-og-tog?tabell=439538 https://www.ssb.no/transport-og-reiseliv/artikler-og-publikasjoner/mindre-utslipp-fra-veitrafikk-fly-og-tog?tabell=439538 https://doi.org/10.1080/09613218.2019.1691486 https://doi.org/10.1016/j.enbuild.2019.109569 acta polytechnica ctu proceedings 38:151–161, 2022 1 introduction 1.1 energy refurbishments 1.2 transport emissions and buildings 1.3 the goal of the study 2 materials and methods 2.1 case study – bærum municipality 2.1.1 background 2.1.2 institutional b-level buildings 2.2 adapting the abc-framework 2.2.1 building emissions 2.2.2 transport emissions 2.2.3 calculations 3 results 3.1 first 30 years-low carbon mix 3.2 remaining 30 years-low carbon mix 3.3 sensitivity analysis 4 discussion 4.1 expanding the scope from building to urban scale 4.2 potential barriers for implementation 5 conclusion references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0546 acta polytechnica ctu proceedings 38:546–552, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague implementation of positive energy district concepts and energy master plans for decarbonization of districts matthias haasea, ∗, daniela baerb a zurich university of applied sciences, department of life sciences and facility management, am gruental, 8820 waedenswil, switzerland b sintef, community, hogskoleringen 7b, 7465 trondheim, norway ∗ corresponding author: matthias.haase@zhaw.ch abstract. in order to achieve a holistic approach to community energy planning for neighbourhoods and districts, it is crucial to provide planners, decision makers, and stakeholders with the necessary methods and instruments. however, there is a research gap in terms of planning and implementation strategies and models. to address this gap, our research used literature and document analysis as well as qualitative interviews to identify implementation models and energy supply options for positive energy districts (peds), and to determine which market actors are needed for peds. we also discussed the consequences of scaling up the ped concept. our analysis highlights the importance of integrated energy planning, which is critical for reducing energy consumption, securing the location of energy infrastructure (generation, distribution, storage), and achieving long-term sustainable development and climate neutrality. therefore, understanding the different dimensions of sustainable development in combination with energy supply and consumption is more important than ever for planning and realizing settlements. keywords: positive energy districts (ped), energy master planning (emp), implementation plans. 1. introduction as policy makers set increasingly ambitious energyrelated building and community requirements and standards based on the sustainable development goals of the united nations (un) [1], climate change presents challenges to achieving these goals. the concept of energy master planning (emp) – a roadmap for planning energy efficiency and grid optimization – can facilitate better planning and implementation to meet these goals. stakeholders at all levels, from nations to communities, face the challenge of reducing greenhouse gas emissions to meet the goals of the paris agreement. a promising approach to reducing energy demand, increasing efficiency, and lowering the carbon footprint is through bottom-up approaches to energy planning at the neighborhood level [2]. cities and communities play a crucial role in achieving global sustainability goals as they are the main source of emissions and have the power to implement global goals at the local level while also considering site-specific demands and settings. 1.1. energy master planning (emp) holistic solutions for meeting the heating, cooling, and power needs of energy communities on a district scale can result in significant energy savings, reduced emissions, and increased energy security. while there is considerable literature available on energy master planning (emp) at the neighborhood and district level, including guidance and assessment tools for campuses [3], existing guidance and tools do not fully address the challenges of emp. energy planning involves determining the optimal mix of energy sources to meet a given energy demand, spanning over the lifetime of a neighborhood. emp comprises six steps, as illustrated in figure 1, starting from goal-setting during the early strategic long-term planning phase and continuing through the operational phase characterized by measurement and verification activities. emp focuses on the tactical mid-term planning and implementation phase, encompassing assessment, scenario building, planning, and implementation [4]. the multi-scale aspect (temporal and geographical) of a holistic approach to neighborhood and district energy planning poses major difficulties, as it requires consideration of both quantitative (economic, technical) and qualitative (environmental impact, social criteria) criteria [5]. to apply this approach and provide necessary methods and instruments for master planners, decision makers, and stakeholders, it is crucial to identify and frame the constraints that bound the options towards an optimized energy master planning solution [6]. existing guidance on master planning emphasizes the critical first step of identifying and establishing project goals [7]. sharp et al. (2020) analyzed emp in several countries and identified the importance of a thorough understanding of local and regional goals and constraints in achieving successful emp [6]. as 546 https://doi.org/10.14311/app.2022.38.0546 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 implementation of positive energy district concepts . . . figure 1. energy master planning process [4]. more countries aim to improve the efficiency, environmental impact, and resilience of buildings and neighborhoods, early and comprehensive energy master planning on the neighborhood and district level becomes increasingly important [4]. however, various constraints can hinder the full potential of decarbonization. in haase and baer (2021), we identified differences in the implementation of positive energy districts (peds) in two distinct cases of decarbonized district solutions, and noted a research gap in planning and implementation strategies and models [8]. 1.2. positive energy districts and emp the positive energy district (ped) concept is rapidly evolving as an integral part of the strategic energy transition (set) plan for european union member states. while a precise definition is still emerging, the idea of creating energy production plants in built environments that generate more energy than the surrounding neighbourhood consumes is a promising way to advance the clean energy transition and achieve climate neutrality through a holistic approach to reducing energy demand and promoting renewable energy. positive energy districts are based on the fundamental principle of creating an area capable of producing more energy than it consumes over the course of a year, with the flexibility to adapt to changes in the energy market [9]. the joint programme of urban europe envisions positive energy districts as “energy-efficient and energy-flexible urban areas or groups of connected buildings that produce net-zero greenhouse gas emissions and actively manage an annual local or regional surplus production of renewable energy. achieving these goals requires integrating different systems and infrastructures, fostering interaction between buildings, users, and regional energy, mobility, and ict systems, and ensuring the energy supply while promoting social, economic, and environmental sustainability for all stakeholders” [10]. in this sense, the ped provides a natural area of the application of emp concepts. the process manager for emp works with municipal departments and external stakeholders to implement and monitor measures in the plans, evaluate results, and communicate feedback. it is important to note that a local energy planning process should be adapted to suit the context and needs of the local community. the local utility is naturally a major player as it traditionally managed the local energy grids. and the process manager has to understand the different implementation models for local energy supply and how they interact with the ped goal. it lacks a constraints analysis to clearly identify energy supply options for peds. 1.3. issues in energy supply (utility model) in recent years the understanding of the role of utilities has changed dramatically. while in the early 1980s utilities were comparably simple heating and cooling supply schemes in existing or new built building clusters (utility services or energy supply contracting) which included supply components on the supply side, and never considered to refurbish the user side. the supply schemes usually had a highly efficient supply technology such as cooling and heating plants (chp), gas-, oilor coal fired and/or biomass supply boilers for hot water or steam and electricity. each of these efficient technologies was backed up by “normal” heating boilers (hot water or steam). the power was usually fed into the power grid (no direct retail to end users) while the heating was distributed by hot water or steam grids to the end users. end users paid a heating price per kwh consumed and a heating load price covering investment costs (at least partially). the end users have been responsible for the demand side: the house station often equipped with a heat exchanger or a steam or hot water exchanger, a domestic hot-water station and a control system including a heating meter for the measurement of the consumed heating energy. with an increasing number of requirements on the national level such as primary or end energy targets for buildings, regulation to foster renewables, partly and temporarily overloaded grids have added multiple complexities which cannot be handled by the utility model alone. in the eu carbon footprint targets 547 matthias haase, daniela baer acta polytechnica ctu proceedings are in place and primary energy (pe) targets for buildings. the latter pe targets can in most cases only be responded positively by combining demand and supply side measures. this means, that a building which is connected to a supply grid with a biomass boiler (low primary energy factor) can even reduce his efforts on the level of the building insulation to achieve the required pe target. this option however is shortsighted, as these buildings will consume large amount of valuable biomass and has consequences for a larger rollout of the concept which is still not well understood. 2. research question building on the identified research gaps we ask three research questions: (i) what are the characteristics of implementation models towards positive energy districts (peds)? (ii) which energy supply options exist and which are needed for peds? (iii) what are the consequences for a larger rollout of the concept? 3. method we describe the characteristics of models for the energy supply towards positive energy districts (peds) by analysing different recent developments of peds. the work is based on literature and document analysis adapted from the iterative model proposed by bocken et al. [11]. in addition, we conducted 34 interviews with focus groups in four different european countries to understand the role of the different stakeholders. the method was based on structured interviews with different stakeholders from practice, policy and technology provider background. 4. results in today’s resource-constrained environment, and with the emergence of “positive energy districts” concepts, local stakeholders are looking for creative ways to drive additional efficiencies in energy use and reduce associated costs. however, large, coordinated efforts are needed to gain synergy between different energy initiatives and future planned projects to maximize energy use and cost reduction. 4.1. characteristics of new energy supply concepts the transformation of today’s electric power sector to a more sustainable energy production based on renewable energies will change the structure of the industry [12]. in this transformation towards a smart energy system interaction between sectors and technologies the main stakeholders as listed above (energy service providers; utilities) will face new challenges in their traditional way of doing business. therefore, adapting their business models to remain competitive is seen as an important step. when looking at the business model literature reveals that there exist basically two possibilities: (i) ownership of renewable energy assets [13] (ii) utilities need to develop from commodity providers to energy service providers [14, 15]. according to this idea utilities should evolve to comprehensive energy-solutions providers for residential and commercial customers to create new sources of revenues. in the energy sector a few characteristics to implement new energy supply concepts are worth noting (see also [11]): • highly resilient community energy system considering buildings, supply, distribution and storage as basic components which have to match together. • target value approach primary energy, carbon footprint target-oriented system approach for the complete community considering incoming and outgoing energy supply streams • understanding each building as a prosumer each building is understood to consume, produce and store energy: pv on the roof, detached chps in complex grid hub structures may be producing energy. • sectoral coupling power produced and not consumed onsite can be attractive to be stored to operate e-vehicle fleets on campuses or delivery services and likewise to match mobility and energy systems. • consequent use of transfer energy systems the supply will consider supply systems which are able to use multiple fuels in the nearer future. one of the important components is chp or fuel cell which are market ready usually operated with natural gas but will have to be transferred to system immanent synthesized or ptf based fuels such as hydrogen or synth gas. • grid-friendly incentivized business model stabile mid and high-tension systems require gridfriendly local communities which means that the input from the community into the grid and vice versa needs to be controlled to keep loads and frequencies balanced in the regional and national grids. this requires business models which provide direct (subsidies, feed in tariffs) and indirect (legal barriers) incentives for escos and utilities to operate the energy system grid friendly. • no more supply-only-concepts buildings are actively integrated in the energy concept following the idea, that the resilience and dependence of imported fuels (on national but also local level of the community) can only be reduced if building follow challenging building standards even far beyond the national minimum requirements. not touching buildings in an emp business model is obsolete 548 vol. 38/2022 implementation of positive energy district concepts . . . from the macroand microeconomic standpoint and also from the position of preserving the asset values in the future. • innovative instant trade energy distribution systems besides the known power, heating, cooling grids in communities it is necessary to provide open direct marketing structures for locally produced power. vienna utilities are one of the front runners to use block chain technologies to distribute energy packages from house to house and apartment to apartment without involving local grid resources which provides first experience-based input [16]. in the us, the brooklyn blockchain system has been identified as one of leading instant trade systems [17]. • remuneration systems the remuneration system is designed in a way that incentivizes the energy service provider to provide the energy in the simplest way to the end consumer, to focus on the energy target for the single building and the complete community and to integrate and distribute normal supply solutions with detached building based renewable power production in a way which is “grid-friendly” at the time given. also add-ons such as coupling with the mobility sector, by providing quick charging stations for e-mobility are remunerated separately. the utility or esco respectively does not increase the margin by selling more energy to the costumer but by keeping the overall target and the energy mix matching with this demand. • user-centric besides the before-mentioned remuneration systems and missing performance monitoring, a key factor is the user behavior: any kind of energetic measure can be countered by obvious or unknowingly misconduct of the users. socioeconomic analysis is still not very common in energy science but the few interviews with users conducted show that users are not capable to operate their apartments sufficiently. this is not only the case for end users in their private homes but also for skilled staff in hospitals and data centers. the assumption “building automation will fix it” turns out not to work. this sets the requirement for the new implementation models to provide simple to operate, intuitive end users hand holds to operate and also control the consumption directly. a few examples in modern housing complexes exist already providing end users, both private and professionals simple app based smart building tools to control the power and heating/cooling consumption in their buildings. 4.2. emerging models energy supply for the buildings stands for the supply of both electric and thermal energy. most of the energy use of buildings is related to space heating (and cooling), therefore the needed thermal energy can be extracted directly from a district heating (and cooling) network (dh(c)) or provided by a heat pump (hp) or even directly converted using electric heaters/coolers. energy supply companies are responsible for supplying buildings and district with their needs in terms of electric and thermal energy, however still most of the provided energy is mostly electric energy. classical electricity supply utilities are suffering nowadays from what is called the “spiral death”, meaning a strong decrease in the number of energy utility customers. the quitting customers are served by new market actors operating with new business models, centered on renewable energy and energy efficiency technologies. the electric energy system contains the physical infrastructure (generation, transport, distribution and use together with their components) and an organized electricity market based on different marketplaces. the market consists mainly of the following actors [18]: • electricity generator, who generates electricity and sell it to the energy suppliers. • electricity suppliers, who purchase the electricity from the generators and sell it to consumers. • transmission system operators (tso), who are responsible for transporting electricity for long distance and ensuring grid stability and reliability by real time dispatch. • distribution network operators (dso), who are responsible for delivering electricity to the consumers and measuring the consumption. • regulators, who set the market rules and oversee the functioning of the market. in the eu countries some energy supply companies operate transnationally, however still a large part of the worlds’ electric energy supply is either based on fossil fuels like coal, gas, and oil or nuclear energy. the production, transmission and distribution of electricity accounts for the largest share of the world’s anthropogenic greenhouse-gas emissions, while the use of emission-free nuclear energy comprises serious security risks and unsolved problems of hazardous waste, therefore the role of renewable energies as the most important instrument to mitigate climate change and reduce negative effects of energy production is increasing. despite the fact that utilities (with national or transnational activities) still have a dominant position they are confronted with disruptions of their current way of doing business and face the challenge to develop new business models for electricity generation from distributed and highly intermittent renewable sources. electrical energy sector is undergoing a continuous process of transformation where a fundamental shift of energy supply towards renewable, co2 neutral energies is taking place, together with a decentralization and digitalization. the classical structure of the electrical energy industry that emerged after the liberalization of the electricity and gas markets 549 matthias haase, daniela baer acta polytechnica ctu proceedings in europe including established business models, is subject to disruptive and massive changes. 4.3. the role of the public the majority of implementation models for peds involve the public sector to some degree, and in many cases the public sector has partial or full ownership of the project. the degree to which the public sector is involved is determined in part by how much it may wish to steer a district energy project towards a variety of local objectives. ped implementation models that are replicable and scalable both technically and financially at the neighborhood, city and national level are key to the acceleration of district energy production? (1.) the “wholly public” implementation model is the most common globally. the public sector, in its role as local authority or public utility, has full ownership of the system, which allows it to have complete control of the project and makes it possible to deliver broader social objectives, such as environmental outcomes and the alleviation of fuel poverty through tariff control. (2.) implementation strategies that focus on “hybrid public and private” energy supply have a rate of return that will attract the private sector, but the public sector is still willing to invest in the project and retain some control. these models can include: (a) a public and private joint venture where investment is provided by both parties that are creating a district energy company, or where the public and private sector finance different assets in the district energy system (e.g. production of heat/cooling versus transmission and distribution); (b) a concession contract where the public sector is involved in the design and development of a project, which is then developed, financed and operated by the private sector, and the city usually has the option to buy back the project in the future; and (c) a community-owned not-for-profit or cooperative business model where a municipality can establish a district energy system as a mutual, community-owned not-for-profit or cooperative. in this model, the local authority takes on a lot of risk initially in development and if it underwrites any finance to the project. (3.) “private” implementation models are pursued where there is a high rate of return for the private sector and require limited public sector support. they are developed as a wholly privately owned special purpose vehicle but may benefit from guaranteed demand from the public sector or a subsidy or local incentives. few cases are developing “private” models as the majority district energy model, also because there are a number of barriers to be lifted. 5. discussion 5.1. implementation models the electric power sector’s value streams are heavily influenced by the regulatory and policy frameworks that govern it. electricity distribution companies’ revenues are regulated by government-appointed regulatory commissions, which can impact the viability of distributed renewable energy businesses (ders) in distribution networks. wholesale electricity markets are subject to market rules established by central authorities like independent system operators (isos) or regional transmission operators (rtos), which are monitored and regulated by energy regulatory commissions (ferc). der business models looking to sell services in these markets must comply with established market rules and regulations. furthermore, the electric power sector is subject to significant national and eu policy support, including subsidies and favorable rules for technologies like solar-based energy generation. understanding these policy and regulatory interdependencies is critical for sustainable business development. community ownership is often seen as a source of income that can be locally controlled, making these investments more socially acceptable as they help develop local supply ownership and prevent value leakage from the local economy. financing renewable energy technologies is a crucial factor for both microgeneration, where infrastructure investment costs are a barrier, and large-scale renewable energy technologies, where upfront costs are often significant. alternative financing sources, like energy cooperatives or crowdfunding platforms, can help mitigate financial risk. in addition to these models, other promising business models include novelty-centered, lock-in-centered, complementarities-centered, and efficiency-centered models [19]. 5.2. supply options for peds the emp must implement goals and constraints on different levels (national, municipal, neighbourhood) and phases. the measurement and verification depend on a number of criteria: • cost-effectiveness of community projects; • economic decision-making criteria: ▷ life-cycle cost calculation (lcc) for emp, ▷ multiple benefits (bankable lcc on the building level), ▷ bankability and risk mitigation of multiple benefits, ▷ cost effectivenes; • investment costs and capital expenditures; 550 vol. 38/2022 implementation of positive energy district concepts . . . • determination of technical concept and investment costs: ▷ gathering of accurate investment costs, ▷ developing detailed energy demand and supply scenario by simulation, ▷ specific risks in the calculation of investment costs; • optimization of investment cost. 5.3. rollout with increasing complexity of the energy supply in building clusters, the partition of energy service companies (esco) of the total market is steadily increasing. today escos and a few innovative utilities are able to provide highly complex energy services, including generation, distribution, storage, selling, m&v combined with demand side measures like refurbishment of buildings, distribution grids and other demand side activities. the business scheme here usually is the energy supply contracting which delivers demand and supply side measures for a fixed investment cost-based price per kw and a price for the consumed kwh of energy. the ownership of all investments, except those in buildings remains for the duration of the contract (5–20 years) with the esco/utility. in recent years, the energy savings performance contracting has been developed into a business model which is not only able to tap energy efficiency potentials in buildings (hvac and thermal envelope) but also provide complex supply, distribution and storage concepts with chp, pv, biomass and heat pumps. here, the remuneration is based on the energy savings and other life cycle cost savings provided by the energy service company. but also other value streams are important to notice: • energy savings; • avoided maintenance and repair costs; • operation cost reduction; • insurance costs; • building comfort and green neighbourhood value; • risks and de-risking methods and tools; • key risk indicators (kri) in general; • kri in emp for building clusters in particular. 6. conclusions the characteristics of implementation models towards positive energy districts (peds) were collected and which energy supply options exist, and which are needed for peds was analysed. the results show a clear trend towards community models that concentrate and optimize value streams within the community. from the analysis of the results, the conclusions are that integrated energy planning is more important than ever. because this will answer the question which energy supply options exist, and which are needed for peds. the understanding of the advanced emp value stream is providing a different mindset which focuses on the needs of the building owners and the community developer. a few cornerstones of the advanced value stream for peds are high resilient community energy system considering buildings, supply, distribution and storage as basic components which have to match together. the main aspects of energy supply models were identified as important to be able to define value streams as basis for successful implementation of decarbonization projects. publicly owned district heating systems provide the potential for a larger rollout of decarbonization. associated value streams were discussed and point to the need for a shift from energy and ghg emission savings to increased comfort on neighborhood level and risk mitigation. to plan and develop settlements that contribute significantly to reducing energy consumption and secure the location of energy infrastructure (generation, distribution, storage), as well as achieve long-term sustainable development and climate neutrality, it is essential to understand the different dimensions of sustainable development in combination with energy supply and consumption. the implementation of energy master plans (emps) requires a range of mandatory and optional services provided to the public owner of a building cluster. acknowledgements this work is related to iea ebc annex 73, annex 75 and annex 83. the discussion among experts is highly appreciated. references [1] un. transforming our world: the 2030 agenda for sustainable development, united nations, 2015. a/res/70/1, [2022-09-22]. https://sustainabledevelopment.un.org/ [2] m. david, s. schönborn. bottom-up energy transition narratives: linking the global with the local? a comparison of three german renewable co-ops. sustainability 10(4), 2018. https://doi.org/10.3390/su10040924 [3] z. huang, h. yu, z. peng, m. zhao. methods and tools for community energy planning: a review. renewable and sustainable energy reviews 42:1335–1348, 2015. https://doi.org/10.1016/j.rser.2014.11.042 [4] m. haase, d. baer. constraints, stakeholders, and framing goals in energy master planning between neighborhood and district. in a. bisello, d. vettorato, d. ludlow, c. baranzelli (eds.), smart and sustainable planning for cities and regions, pp. 3–13. springer international publishing, cham, 2021. https://doi.org/10.1007/978-3-030-57764-3_1 [5] j. schiefelbein, c. slotterback, j.-p. petersen, et al. implementation of energy strategies in communities – results within the context of iea annex 63. in 30th international conference on efficiency, cost, 551 https://sustainabledevelopment.un.org/ https://doi.org/10.3390/su10040924 https://doi.org/10.1016/j.rser.2014.11.042 https://doi.org/10.1007/978-3-030-57764-3_1 matthias haase, daniela baer acta polytechnica ctu proceedings optimization, simulation and environmental impact of energy systems, ecos 2017, pp. 1–10. san diego, us, 2017. 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[19] m. haase, e. motoasca. characteristics of business models for innovation clusters in decarbonation projects. acta polytechnica ctu proceedings 38:417–423, 2022. central europe towards sustainable building 2022 (cesb22). https://doi.org/10.14311/app.2022.38.0417 552 https://www.techstreet.com/ashrae/standards/or-20-020-energy-master-planning-identifying-framing-constraints-that-scope-your-technology-options?product_id=2113219 https://www.techstreet.com/ashrae/standards/or-20-020-energy-master-planning-identifying-framing-constraints-that-scope-your-technology-options?product_id=2113219 https://www.techstreet.com/ashrae/standards/or-20-020-energy-master-planning-identifying-framing-constraints-that-scope-your-technology-options?product_id=2113219 https://www.techstreet.com/ashrae/standards/or-20-020-energy-master-planning-identifying-framing-constraints-that-scope-your-technology-options?product_id=2113219 https://doi.org/10.1016/j.enbuild.2017.08.074 https://doi.org/doi/10.2777/04888 https://jpi-urbaneurope.eu/ped/ https://doi.org/10.1016/j.jclepro.2013.11.039 https://doi.org/10.1016/j.rser.2012.01.072 https://doi.org/10.2172/924651 https://www.wien.gv.at/stadtentwicklung/energie/pdf/energierahmenstrategie-2030-en.pdf https://www.wien.gv.at/stadtentwicklung/energie/pdf/energierahmenstrategie-2030-en.pdf https://doi.org/10.1016/j.rser.2018.10.014 https://doi.org/10.14311/app.2022.38.0417 acta polytechnica ctu proceedings 38:546–552, 2022 1 introduction 1.1 energy master planning (emp) 1.2 positive energy districts and emp 1.3 issues in energy supply (utility model) 2 research question 3 method 4 results 4.1 characteristics of new energy supply concepts 4.2 emerging models 4.3 the role of the public 5 discussion 5.1 implementation models 5.2 supply options for peds 5.3 rollout 6 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0025 acta polytechnica ctu proceedings 38:25–30, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague quantifying the impact of external and internal factors and their interactions on thermal load behaviour of a building christoph matschia, ∗, isabell nemethb a hochschule ansbach, campus feuchtwangen, an der hochschule 1, 91555 feuchtwangen, germany b technische hochschule rosenheim, fakultät für angewandte naturund geisteswissenschaften, hochschulstraße 1, 83024 rosenheim, germany ∗ corresponding author: christoph.matschi@hs-ansbach.de abstract. for the energy-efficient design of district heating networks, knowledge about the neighborhood heat load behavior, through heating load profiles in high temporal and spatial resolution, is crucial. due to the high effort required for transient calculations, a less complex method is needed at the neighborhood level. for this reason, a method is developed, which identifies the relevant parameters influencing the building heating load behavior. taking these parameters into account, a simple method for heating load profiling is developed using a machine learning algorithm. for this purpose, a parameter study is conducted using dynamic thermal building simulation software. different parameters influencing the building heating load behavior are varied. to determine the strength of the influence of the individual parameters on the building heating load, to check whether the influence of the parameters is constant or varies over the year and whether parameters are missing here, the results of the parameter study are evaluated statistically. first results show promising results in the detection of the significant parameters, for the creation of a model based on a machine learning algorithm, and the possibility of quantifying their impact on building heating load behaviour. keywords: energy-efficient building design, sector coupling, thermal load behaviour, standardised and parameterised thermal load curves. 1. introduction globally, the building sector accounts for a large share of total energy demand and greenhouse gas emissions. in the eu, the building sector is estimated to be responsible for 40 % of total energy demand and about 36 % of greenhouse gas emissions [1]. in the united states, the building sector accounts for 38.9 % of energy demand, with heating, cooling, and ventilation of buildings already accounting for 34.8 % [2]. these facts highlight the great potential for energy and greenhouse gas savings in the building sector. in order to exploit this potential, not only an energetically optimized building envelope and an optimization of energy generation and transmission technology are required, but also a targeted establishment of technical building equipment, since local and district heating networks are being expanded more and more worldwide. most research on the optimization of local and district heating networks is still based on energy generation and only a few consider the consumer side [3]. however, a suitable prediction of the load behavior on the load side is a fundamental prerequisite for the development and optimization of local and district heating networks. in particular, for an energy-efficient design of the networks and their components, the knowledge of the temporally high-resolution course of the load as well as the simultaneity of the heat demand is of central importance. these parameters, the coincidence factor and the heat load curve, are currently determined using static methods and assigned safety factors due to the resulting uncertainties [4]. as a result, systems are not operated in terms of optimal energy efficiency. in practice, local and district heating networks are designed and optimized on the basis of previous findings. historical data from other local and district heating networks are used to predict the load behavior of an area supplied. since different areas may differ greatly in their underlying conditions, the inherent data characteristics must be adapted for the new area. how well this transfer fits, depends heavily on the experience of the planner [5]. in order to enable a differentiated and quantitative prediction of the heat load demand for the planning and optimization of local and district heating networks, some research has been done in the last years [3]. on the one hand, so-called top-down models have been developed, which are based on historical data of existing local and district heating networks. however, these models have proven to be too coarse and too general for network planning and optimization. bottom-up models, on the other hand, which consider individual nodes in detail as a basis for operation, for example, can be divided into four categories: empirical models, statistical models, physical-statistical hybrid models, and engineering models. while models in the first three categories have a high level of detail, 25 https://doi.org/10.14311/app.2022.38.0025 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en christoph matschi, isabell nemeth acta polytechnica ctu proceedings stage experimental plan climate try 2 try 4 try 14 wind city centre in the countryside sea building tightness tight medium tight leaky window size 10 % 100 % 200 % orientation north south west east table 1. experimental plan. compared to top-down models, they still have disadvantages, especially in the early stages of planning, because they are also based on historical data and are therefore not transferable. it is not possible to implement effects that are not included in the observed data in these models. the engineering models allow the implementation of all required physical effects and user behavior and are based on thermal building simulations. this makes them highly adaptive and also quite accurate in predicting heat demand. however, this approach is very complex and thus expensive to create, since each unit of a neighborhood to be supplied must be simulated individually. to facilitate this, general load profiles can be developed for specific building types, e.g. office buildings, residential buildings, etc. while this simplification speeds up the development of the heat load profiles, it comes at the expense of accuracy as it relates to average building data. for a more detailed overview of heat load profile modeling, the reader is referred to [3], where the authors provide a comprehensive review of current assessment methods. following the latter approach, the goal of this research is to produce heat demand profiles with reasonable accuracy and the advantage of a less timeconsuming setup, compared to methods, where all units are assessed individually. for this purpose, a method is developed to identify the relevant parameters that influence the heating load of a building. on the basis of these identified influencing factors, a simple method for the creation of heating load profiles will be developed with the help of a machine learning algorithm. in this work, the method for identifying the influencing variables is tested. 2. method according to the need to gain insight into the load behavior of the building without a complex transient simulation, the method aims at reducing the building analysis to the crucial parameters of the heat load profiles. in contrast to a coarse simulation for the whole building with only one node, in this work a method is developed and tested making an identification of the most important parameters is possible. with the help of these identified parameters, a method will be developed at a later stage which, with the help of machine learning algorithms, can quickly generate precise heating load profiles. the approach is induced by the task to provide profiles for a long-term simulation and a linear optimization of a district heating network and its heat generators and to develop its future-oriented energy concept. 2.1. principle procedure the individual simulations of the test series for this work are very time-consuming. for this reason, only one building and only one user profile are used to test the method. the basis of the investigations is the small residential building (efh_klein) from the research report “entwicklung einer datenbank mit modellgebäuden für energiebezogene untersuchungen, insbesondere der wirtschaftlichkeit” [6]. using this building, the variables generally considered as relevant, such as storage mass, window size, building orientation, etc., are varied and simulated with the help of a parameter study. subsequently, the simulation results from the parameter study are statistically evaluated with the help of an anova (analysis of variance). in order to find out which parameters have a significant influence on the heating load behavior of a building, the small residential building from the research report “entwicklung einer datenbank mit modellgebäuden für energiebezogene untersuchungen, insbesondere der wirtschaftlichkeit” [6] was first implemented in the dynamic thermal building simulation software ida ice [7]. in this first step, a residential use from din v 18599-10 [8] was assumed as the user profile and user behavior and was not varied further. thus, the results are based on a small residential building and may differ for a large building and/or a different use to these. the varying parameters were changed in a two to four step experimental design. the parameters climate, taken from din 4710 [9], wind, building tightness and window size were varied in a three-stage test plan. the parameter building orientation was varied in a four-stage test plan, see table 1. the building insulation and storage mass were varied in a two-stage experimental design. for the building insulation, the tabula database [10] was used as a basis for the varying component structures. the uvalues and the corresponding storage mass are shown in table 2. all parameters are varied and simulated using the experimental plan table 1. 26 vol. 38/2022 quantifying the impact of external and internal factors . . . heavy lightu-value storage mass [w/m2k] [kj/m2k] exterior wall new construction 0.136 484 24 exterior wall not refurbished 1.52 352 16.6 interior wall 74 16.1 false ceiling 370 176 base plate new construction 0.21 531 348 base plate not refurbished 0.95 495 215 roof new construction 0.14 393 29 roof not refurbished 1.81 281 18.1 table 2. characteristic values of the building components. for these investigations, as mentioned above, anovas were performed to find out to what extent the influencing parameters and their interaction have an influence on the building heating load. for this purpose, the average heating load was determined for each month of the year and used as a dependent variable. the varied influence parameters from the test plan form the independent / explanatory variables in the analysis. thus, 12 anovas were performed, one for each month. subsequently, for each independent variable and interaction among the variables, the partial η2 was used to determine the influence strength on the building heating load within the anova. in a further step it was examined whether the influence strength changes over the year. finally, it was examined, whether the coefficient of determination r2 of the individual anovas changes over the year. 2.2. model set up the software “ida ice” of the software manufacturer equa provides the possibility of parameterized simulations. this makes it comparatively easy to carry out parameter studies with several varying parameters. following the parameterized simulations, the generated data can be relatively easily merged in excel for further analysis. at the beginning of a series of tests, the parameters within this series are permanently set according to the experimental setup in table 1. for example, before a parameter simulation, it is determined whether it is a heavy or light building. furthermore, the quality of the building envelope is defined. here it is distinguished, whether it is a “not refurbished” building envelope or whether it is a “new” building. the components for the category “not refurbished” would correspond to buildings of the age class 1816 to 1918, the components of the category “new construction” to the age classes 2016 to today. the characteristic values of the building components are shown in table 2. the parameters to be varied are defined in the ida ice software at the beginning and automatically varied during the parameter simulations. for the climate, the test reference years try 2, try 4 and try 14 of din 4710 [9] were used. the test reference years contain hourly values of, for example, outdoor temperature, solar radiation, etc. for more detailed information see din 4710 [9]. for the wind profiles (wind), the ashrae profiles “in the countryside”, “sea” as well as “city centre” stored in the software ida ice were used. the n50 air change rate was used to characterize the building tightness. this defines how often the air exchanges in one hour at a pressure difference of 50 pa. in the “tight” category, an n50 value of 0.35/h was assumed, in the “medium tightness” category an n50 value of 3.675/h was assumed, and in the “leaky” category an n50 value of 7/h was assumed. window sizes were simulated at 10 %, 100 % and 200 % starting from the initial model. from the research report “entwicklung einer datenbank mit modellgebäuden für energiebezogene untersuchungen, insbesondere der wirtschaftlichkeit” [6], the small residential building has a volume of 465 m3 in its basic configuration and has 26.5 m2 of window area, a usable floor area of 148.8 m2 and an a/v ratio of 0.98. the building orientation was varied in four steps starting from a south orientation to a west, north and east orientation. 3. results the following section presents the results for testing the method of quantifying the effects of external and internal factors and their interactions on the thermal load behavior of a building and interprets the results statistical analysis on the parameter studies of selected influencing factors and their interactions, as well as the change in model quality over the course of the year. 3.1. difference in the strength of influence of individual parameters and their interactions over the course of the year the first step was to investigate the extent to which the influence of the individual independent variables and their interactions change over the year. due to the abundance of independent variables and especially 27 christoph matschi, isabell nemeth acta polytechnica ctu proceedings data series anova-model parameter (variable and interaction) 1 wind 2 building tightness 3 climate * building tightness 4 window size * wind 5 wind * building tightness 6 wind * average u-value building 7 climate * window size * wind 8 climate * wind * building tightness 9 window size * orientation * average u-value building 10 climate * window size * wind * building tightness table 3. selected data series with model parameters. figure 1. influence of data series 1 and 2 over the year. figure 2. influence of data series 4 and 6 over the year. their interactions, which leads to a total of 1585 model parameters (variables and their interactions) within an anova, the 10 most important data series are shown as examples. the presented data series with their model parameters are listed in table 3. in order to quantify the strength of influence or effect of the individual variables and their interactions on the overall model, the partial eta-squared η2, which results from the quotient of the explainable sum of squares to the total sum of squares, was calculated figure 3. influence of data series 5 and 8 over the year. figure 4. influence of data series 3, 7 and 9 over the year. for each data series. according to cohen [11], a η2 of 0.01 indicates a small effect, whereas a η2 of 0.14 and higher indicates a large effect. the following figures show the influence of the individual variables and their interactions over the year. as shown in figures 1, 2, 3 and 4, the influence of the individual variables and their interactions change over the year, except for data series 2, where the influence on the heating load behavior remains relatively constant over the year. with the data series and the variables and interactions contained therein 1 as well 28 vol. 38/2022 quantifying the impact of external and internal factors . . . figure 5. coefficient of determination of the anovas over one year. as 3 to 10 it is to be recognized that their influence of the winter months to the summer months decreases and in the connection again rises. if we look at the variable in data series 2 in combination with the variable climate, which corresponds to data series 3, we can see that the building tightness, which corresponds to the variable in data series 2, also loses importance from the winter months to the summer months and vice versa gains importance again. it seems to be possible to determine the influence of individual parameters and their interactions with the help of this method. in addition, this analysis indicates that the significance of the influencing variables and their interactions change over the year. 3.2. coefficient of determination of the anovas next, it was checked whether the coefficient of determination r2 of the anovas remains constant over the year or whether there are differences. for this purpose, the coefficient of determination r2 was calculated for each anova and compared with each other. as shown in figure 5, the quality r2 of the anovas decreases from the winter months to the summer months and then increases again. in the winter months the r2 is about 0.91 which indicates a very good predictability of the building heating load by the anova. in the summer months, the r2 decreases to about 0.6, indicating only moderate predictability. overall, this decrease in r2 from the winter months to the summer months indicates that explanatory variables are missing here to be able to predict the building heating load well in summer. this means that more variables have to be included and not all explanatory variables are included in the regression model to predict the building heating load behavior in summer. this means that not all variables that have an influence on the heating load behavior of a building have been identified yet. 4. discussion the results from section 3 show that it is possible to determine the strength of influence of individual parameters and their interactions. furthermore, the analyses show that these vary strongly over the year. also, it seems to be possible to determine with this method whether all important parameters for the prediction of the heating load behavior of buildings have been found. as shown in section 3.2, the decrease of r2 from winter to summer and vice versa indicates that not all important parameters have been included in the analysis yet. in a next step, further analyses will be performed to determine the last missing parameters that have an influence on the building heating load behavior. afterwards a principal component analysis will be performed to check if variables can be combined. this should reduce the input effort for a later heat load prediction model. with the principal component analysis, the influence of the individual parameters on the building heating load model can be determined and quantified even more precisely. the influential parameters will ultimately serve as input variables for a machine learning algorithm, which will then be used to predict the building heating load behavior with similar precision as with a dynamic thermal building simulation, but with considerably less effort. with this algorithm, the heating load behavior of entire neighborhoods should also be precisely presentable without great effort. references [1] european commission. 2009 statistics & market observatory. [2014-07-26], http://ec.europa.eu/energy/observatory/. [2] r. j. de dear, t. akimoto, e. a. arens, et al. progress in thermal comfort research over the last twenty years. indoor air 23(6):442–461, 2013. https://doi.org/10.1111/ina.12046. [3] w. ma, s. fang, g. liu, r. zhou. modeling of district load forecasting for distributed energy system. applied energy 204:181–205, 2017. https://doi.org/10.1016/j.apenergy.2017.07.009. 29 http://ec.europa.eu/energy/observatory/ https://doi.org/10.1111/ina.12046 https://doi.org/10.1016/j.apenergy.2017.07.009 christoph matschi, isabell nemeth acta polytechnica ctu proceedings [4] bwk gleichzeitigkeit – der unterschätzte faktor. [2018-11-21], https://www.rehau.com/download/ 1475518/fachartikel-gleichzeitigkeit.pdf. [5] t. nussbaumer, s. thalmann, a. jenni, j. ködel. planungshandbuch fernwaerme. energieschweiz, 2017. isbn 3-908705-30-4. [6] entwicklung einer datenbank mit modellgebäuden für energiebezogene untersuchungen, insbesondere der wirtschaftlichkeit. dipl.-ing. swen klauß zentrum für umweltbewusstes bauen e.v. verein an der universität kassel prof. dr. anton maas (vorstand) gottschalkstraße 28a 34127 kassel. [7] ida ice 4.8 sp2, equa solutions ag, untermüli 3 6300 zug schweiz. [software]. [8] din v 18599-10:2011-12. [din-standard]. [9] din 4710:2003-01: statistiken meteorologischer daten zur berechnung des energiebedarfs von heizund raumlufttechnischen anlagen in deutschland. [10] tabula. [2021-10-01], https://webtool.building-typology.eu/?c=all#bm. [11] j. cohen. statistical power analysis for the behavioral sciences. taylor and francis, hoboken, 1988. 30 https://www.rehau.com/download/1475518/fachartikel-gleichzeitigkeit.pdf https://www.rehau.com/download/1475518/fachartikel-gleichzeitigkeit.pdf https://webtool.building-typology.eu/?c=all#bm acta polytechnica ctu proceedings 38:25–30, 2022 1 introduction 2 method 2.1 principle procedure 2.2 model set up 3 results 3.1 difference in the strength of influence of individual parameters and their interactions over the course of the year 3.2 coefficient of determination of the anovas 4 discussion references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0649 acta polytechnica ctu proceedings 38:649–655, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague catalogue of construction products containing secondary raw materials from different industries and municipal waste teraza pavlůa, jan peštaa, b, ∗, martin volfa, antonín lupíšeka a university center for energy efficient buildings of technical university in prague, třinecká 1024, 273 43 buštěhrad, czech republic b university of chemical technology in prague, faculty of environmental technology, department of sustainability and product ecology, jankovcova 23, 170 00 prague 7, czech ∗ corresponding author: jan.pesta@cvut.cz abstract. the building industry consumes a large amount of primary raw materials and also contributes significantly to the production of waste. applying circular principles in this field to reduce resource consumption and waste production has been investigated in several projects considering the reuse or recycling of construction and demolition waste. however, consumption of primary raw materials can also be reduced by re-targetting waste from different industries and municipal waste to produce new construction products. thus, opportunities for the recycling of industrial and municipal waste were investigated in this project. the main output is the catalogue, which provides an overview of products with recycled content and secondary materials with the potential to be used in the construction industry such as blast furnace slag, ash, and energy by-products. also, it contains a list of valid requirements for the use of recycled materials under specific conditions of the czech republic. in addition, examples of good practice are presented to break the existing behavioral barriers to the use of secondary raw materials in the czech construction industry. this contribution summarizes the findings in the field of industrial and municipal waste recycling and its further use as secondary raw materials in the construction industry. keywords: waste, circular economy, secondary raw materials, primary raw materials, recycled materials, recycled content. 1. introduction sustainable buildings are based on sustainable resources, which are used for construction products. however, how sustainable are the sources of primary raw materials for the building industry? the amount of materials consumed in the global economy is more than ten times higher than the amount of reused or recycled materials [1]. the difference between these two flows is called the circularity gap and is increasing annually. this is significantly affected due to the construction industry. on the one hand, the construction industry consumes more than 30 % of domestic material consumption in the czech republic [2]. on the other hand, construction and demolition waste represents almost one-third of european waste [3]. in an effort to increase circularity in the european economy, the european commission published the circular economy action plan [4] as one of the supporting strategies for the green deal [5]. both of these documents described construction and building as a key sector for reducing the impact on climate change and exploiting resources. meanwhile, the main regulation for high rates of recycling materials and their utilization in the construction industry is defined by regulation (eu) no. 305/2011 of the european parliament and the council, which focuses on sustainable utilization of natural resources [6]. the development of this effort under the conditions of the czech republic is mainly defined by the waste management plan, which defines the minimum recycling rate of construction waste as 70 % [7]. to support the recycling of demolition and construction waste, many activities were carried out. one of them was the publication of a catalogue of construction products containing recycled materials from construction and demolition waste [8]. however, the circularity of the whole economy can also be improved by re-targetting waste from different sources to industries, where it can be recycled or reused. to improve the re-targeting of resources from different industries and sectors, the second part of the previously mentioned catalog will be published at the end of 2021 [9]. the catalogue of construction products containing secondary raw materials from different industries and municipal waste was prepared as a collection of information from experts, producers, and civil engineers and was discussed at round tables in cooperation with czech authorities. in addition, the prioritization of key waste flows with a high potential for re-targetting to the construction industry was performed considering national statistics of secondary raw materials 649 https://doi.org/10.14311/app.2022.38.0649 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en t. pavlů, j. pešta, m. volf, a. lupíšek acta polytechnica ctu proceedings figure 1. production of secondary raw materials in the czech republic [million tons] [10]. production. this catalogue contributes by summarizing possible uses, risks, and barriers for recycling of waste in the building industry. it will be used by architects, designers, civil engineers, construction contractors, and public and private investors who want to improve the circularity of their projects or find new sources for their constructions and products. in the catalogue, an overview of products with recycled content is presented together with the list of valid requirements on the utilization of recycled materials listed in standards and legislation. moreover, the catalogue provides examples of good practices to decrease suspicions about the use of recycled materials. this contribution summarizes the findings on the possible use of secondary raw materials from different industries and municipal waste in the construction industry. 2. methods the preparation of this catalogue followed after publication of the previous catalogue of construction products containing recycled materials from construction and demolition waste. similarly, the new part is based on the evaluation of national statistics describing waste production. in addition, information about the possible utilization of recycled materials in the construction industry was summarized after discussions and roundtables with experts. the group of experts included producers of industrial waste, civil engineers, and producers of construction products. the outputs of this group were consulted with state officials representing perspective focused on the protection of nature and public health. the discussions were organized in collaboration with the ministry of industry and trade and the czech standardization agency. based on evaluation of national statistics and summarization of possible utilization following key waste flows with the highest potential to be recycled were descripted in this catalogue: slag, ashes and fly ashes, energy gypsum, plastic materials, glass, paper, composite packaging, tires and rubber, wood, and metals. prioritizing key waste flows was followed by a broad literature study of available standards, existing legislative documents, and regulation. the possible use of recycled materials was declared in the catalogue on the basis of communication with companies and good examples from construction practice. the next stage of preparation of catalogue was collecting of feedback from experts, companies, and state officials. finally, the results of the catalog will be published as a freely available pdf document on the web http://www.recyklujmestavby.cz/, which was developed for the first part of the catalogue [9]. 3. results the final version of the catalog will be published at the end of november 2021. however, a summary of the main outcomes is provided in the following chapters. 3.1. production of secondary raw materials secondary raw material can be characterized as a product (including certified products) or recovered waste, which stopped being waste after meeting the recovery regulations. in 2018, the production of secondary raw materials was 22.2 million tuns [10]. this production increased in comparison with the production in 2017. in comparison to the extraction of primary materials in the czech republic, it represents approximately one third of the excavated amount of construction materials (more than 64 million tuns) [12]. based on this, the high potential for saving of primary resources can be assumed. the largest amount of secondary materials is produced from energy processes and reached (including by-products) more than 10.2 million tun in the year 2018. the overview of the flows with the biggest amount is presented in figure 1. the flows 650 http://www.recyklujmestavby.cz/ vol. 38/2022 catalogue of construction products containing secondary raw . . . type of secondary raw material from energy processes production [t] percentage [%] fly ash from combustion of coal 7 322 239 60.0 slag 1 287 574 10.5 fly ash from fluid combustion of coal or a coal with biomass 1 452 175 11.9 fly ash from fluid combustion of biomass 118 783 1.0 fly ash from comb. of biomass (except fluidized bed boiler) 5 707 0.0 sda product 70 380 0.6 energy gypsum 1 952 843 16.0 table 1. production of by products from energy sources in 2017 [11]. specification the main risks to reuse andrecycling possibilities of utilization slag aggregates – grinding and sorting of crystal slag disintegration of slag in contact with water. backfilling, landscaping. as a layer under construction foundations. aggregates for concrete mixtures. prefabricated construction concrete elements. granulated slag – blast furnace slag produced after cooling with water disintegration of granulated slag is slow process. after milling used in the cement production. hydraulic road binder. boiler slag – the solid waste from combustion of solid fuels different composition. admixture for brick production to optimize properties of ceramic material. table 2. possible use of slag in buildings. of secondary materials are more specified in table 1. the most important energy source for the by-products is coal combustion, in which fly ash is produced [11]. 3.2. possible use of secondary raw materials 3.2.1. slag slag with potential to be used in buildings is produced as a byproduct during steel production in thermal processes and combustion. because of these sources, slag is available in a large amount and with a low price. on the other hand, the variability of sources is connected to variability of properties. for example, slag aggregates can disintegrate when in contact with water. based on this, use of slag in practice is limited by suspicion of practitioners. however, this main barrier can be overcome with a proper description of the original source of the slag. other barriers are described in table 2. 3.2.2. ashes and fly ash several types of fly ashes are described in table 3. fly ashes are used as an admixture for cement production. also, it is produced in high amounts and at low cost as an alternative for the cement industry. however, the main barrier is suspicion to this material, which is produced as a byproduct. however, fly ash must comply with the regulations for the declaration of ecotoxicological safety. 3.2.3. energy gypsum energy gypsum is produced as a by-product in the process of wet scrubbing of flue gas using a solution with limestone. the recycling or reuse of gypsum materials and products is a highly recommended way of waste management because landfilling of gypsum materials can be allowed only under specific conditions, which will reduce the potential contact with organic matter leading to the production of hydrogen sulfide. the possible use of energy gypsum is described in table 4. however, the production of gypsum in the future can be reduced due to the closure of thermal power plants. 3.2.4. plastic materials in table 5, the specific use of recycled plastic materials is described. the main barrier to recycling plastic waste is insufficient plastic waste classification. municipal waste can contain tens of types of platic materials and for their use sorting capacity is needed. moreover, the content of toxic chemical substances in plastic materials should be tested in case this material is used on place, where, for example, there is a high probability of contact with human skin. 3.2.5. glass the glass is a material, which can be recycled multiple times. moreover, using waste glass as an admixture leads to a reduction in the melting temperature in 651 t. pavlů, j. pešta, m. volf, a. lupíšek acta polytechnica ctu proceedings specification the main risks to reuse andrecycling possibilities of utilization fly ash – produced in process with high-temperature combustion the utilization is limited by regulation describing technical properties for future use. pozzolanic properties can be reduced during storing in space with high humidity. high difference of properties based on specific sources. replacement of cement. aggregates or filler for concrete and prefabricated concrete elements. production of aerated autoclaved concrete. admixture for brick production. mineral fibers. thermal insulation products for buildings. fluid ash – produced in process with fluid incineration high difference of properties based on specific sources. mortar production. aggregates or filler for concrete. hydraulic road binder. thermal insulation products for buildings. production of aerated autoclaved concrete. admixture for brick production. mineral fibers. table 3. possible use of ash and fly ash in buildings. specification the main risks to reuse andrecycling possibilities of utilization energy gypsum the possible use in gypsum plasterboard is limited by technical regulations that define the gypsum parameters. gypsum production. gypsum binders. cement production. gypsum plasterborard. gypsum fibreboard. table 4. possible use of energy gypsum in buildings. specification the main risks to reuse andrecycling possibilities of utilization recycled polyvinyl chloride (pvc) insufficient sorting of pvc backfilling. landscaping. as a layer under construction foundations. aggregates for concrete mixtures. recycled plastic materials from municipal waste variable composition of municipal waste. high amount of types of plastic materials. products and accessories for garden. covering of terrace, pavement, fences. recycled polyester insufficient sorting of pe geotextiles recycled polystyrene (eps) insufficient sorting of eps recycled eps recycled plastic insulation for cables insufficient sorting of pvc floor covering recycled materials from automobile interiors inseparable mixture of plastic materials and textile materials. acoustic insulation table 5. possible use of plastic materials in buildings. 652 vol. 38/2022 catalogue of construction products containing secondary raw . . . specification the main risks to reuse andrecycling possibilities of utilization flat glass – from deconstruction or as waste from automobile industry an insufficiently sorted batch of glass can be devalued by impurities and different types of glass. window fillings. recycled glass insufficient sorting glass fibers – thermal and acoustic insulations, foam glass. lightweight aggregates for concrete. glass fibers as reinforcement. table 6. possible use of recycled glass in buildings. specification the main risks to reuse andrecycling possibilities of utilization recycled paper insufficient sorting blown cellulose fibers as thermal and acoustic insulation. fibers as reinforcement. table 7. possible use of paper in buildings. specification the main risks to reuse andrecycling possibilities of utilization recycled composite packaging disintegration of slag in contact with water. boards as construction elements, roof covering. table 8. possible use of recycled composite packaging in buildings. specification the main risks to reuse andrecycling possibilities of utilization recycled tires and recycled rubber the health risk should be secured according to the regulations for the use of products containing recycled rubber. coverings of playgrounds. floors for industry, coverings for balconies and terraces. insulation against vibration. table 9. possible use of tires in buildings. specification the main risks to reuse andrecycling possibilities of utilization furniture wood possible contamination with dangerous substances (glues, paint) woodfibre boards and insulations wood from wood industry production of pellets for energyrecovery. wood from municipal waste wood can be damaged by biological processes or infested by pests. composting. energy recovery. table 10. possible use of recycled composite packaging in buildings. 653 t. pavlů, j. pešta, m. volf, a. lupíšek acta polytechnica ctu proceedings specification the main risks to reuse andrecycling possibilities of utilization metal packaging, metal parts of machines and tools, and metal waste from the manufacturing industry insufficient sorting. production of new metal partsand products. table 11. possible use of recycled metals packaging in buildings. batches for the production of glass from primary resources. other risks and possible use are described in table 6. 3.2.6. paper paper material can be contaminated by impurities from municipal waste. however, in case of proper sorting, the paper material has several possibilities for utilization, which are described in table 7. 3.2.7. composite packaging material recycling of composite packaging is limited due to the number of materials used and the technology of their connection. thus, the composite packaging is often just shredded. the possible utilization of this material is described in table 8. 3.2.8. tires and rubbers in table 9, the possible use of tires and rubbers is described. the main barrier to their recycling is the risk that the recycled material contains dangerous substances. therefore, these materials must be tested and fulfil the regulations for their use. 3.2.9. wood taking into account the effort to circularity, wood should be reused or recycled, but, in some cases, incineration with energy recovery can be a more economical way of waste management. however, possible use for recycling or reusing wood waste is described in table 10. the main barrier to the reuse of wood is possible contamination by pests or dangerous substances. 3.2.10. metals the sorting of metals is supported by high value of scrap, which can be easily sold in facility for waste collection. in addition, metals have high potential to be recycled after sorting due to the high cost of the primary resource and the sufficient infrastructure for recycling. the possible use of metal is briefly described in table 11. 4. conclusion the catalogue contributes to increasing the circularity of materials in the czech economy by identifying key material flows with high potential to be used in the building industry. moreover, the catalogue presents the products with recycled content and examples of good practices. furthermore, the possible use of secondary raw materials is described according to the european and czech legislation. for each of the key material flows, risks and barriers in the recycling process were described. one of the main barriers is the suspicion against the use of products with recycled content. nevertheless, the products containing recycled materials must fulfil at least the same regulations on the market as the products made from primary resources. further research is planned on possibilities of how to re-target waste materials and increase circularity. however, research should focus not only on retargetting of materials but also on improving materials, as inputs to the economy, to be recyclable in the future. acknowledgements this work was supported from the grant of specific university research – grant no a1_ftop_2021_003. references [1] the circularity gap summary 2021. [2022-03-03]. https://drive.google.com/file/d/1_dqhx5ztajz_ 4nyp72gjijej_yqchnxw/view [2] zpráva o životním prostředí české republiky 2015. [2022-03-03]. https: //www.mzp.cz/c1257458002f0dc7/cz/news_161129_ zprava_o_zp/$file/zprava_zp_2015_291116.pdf [3] strategy for a sustainable built environment. [2022-0303]. https://www.europarl.europa.eu/legislativetrain/theme-a-european-green-deal/filestrategy-for-a-sustainable-built-environment [4] a new circular economy action plan for a cleaner and more competitive europe. [2022-03-03]. https://eur-lex.europa.eu/legal-content/en/txt/ ?qid=1583933814386&uri=com:2020:98:fin [5] the european green deal. [2022-03-03]. https://eur-lex.europa.eu/legal-content/en/txt/ ?qid=1588580774040&uri=celex:52019dc0640 [6] regulation (eu) no 305/2011 of the european parliament and of the council of 9 march 2011 laying down harmonised conditions for the marketing of construction products and repealing council directive 89/106/eec. [2022-03-03]. https://eur-lex.europa.eu/legal-content/en/txt/ html/?uri=celex:32011r0305&from=cs 654 https://drive.google.com/file/d/1_dqhx5ztajz_4nyp72gjijej_yqchnxw/view https://drive.google.com/file/d/1_dqhx5ztajz_4nyp72gjijej_yqchnxw/view https://www.mzp.cz/c1257458002f0dc7/cz/news_161129_zprava_o_zp/$file/zprava_zp_2015_291116.pdf https://www.mzp.cz/c1257458002f0dc7/cz/news_161129_zprava_o_zp/$file/zprava_zp_2015_291116.pdf https://www.mzp.cz/c1257458002f0dc7/cz/news_161129_zprava_o_zp/$file/zprava_zp_2015_291116.pdf https://www.europarl.europa.eu/legislative-train/theme-a-european-green-deal/file-strategy-for-a-sustainable-built-environment https://www.europarl.europa.eu/legislative-train/theme-a-european-green-deal/file-strategy-for-a-sustainable-built-environment https://www.europarl.europa.eu/legislative-train/theme-a-european-green-deal/file-strategy-for-a-sustainable-built-environment https://eur-lex.europa.eu/legal-content/en/txt/?qid=1583933814386&uri=com:2020:98:fin https://eur-lex.europa.eu/legal-content/en/txt/?qid=1583933814386&uri=com:2020:98:fin https://eur-lex.europa.eu/legal-content/en/txt/?qid=1588580774040&uri=celex:52019dc0640 https://eur-lex.europa.eu/legal-content/en/txt/?qid=1588580774040&uri=celex:52019dc0640 https://eur-lex.europa.eu/legal-content/en/txt/html/?uri=celex:32011r0305&from=cs https://eur-lex.europa.eu/legal-content/en/txt/html/?uri=celex:32011r0305&from=cs vol. 38/2022 catalogue of construction products containing secondary raw . . . [7] plán odpadového hospodářství české republiky pro období 2015–2024, 2014. [2022-03-03]. https://www.mzp.cz/c1257458002f0dc7/cz/poh_cr_ prislusne_dokumenty/%24file/oodp-poh_cr_20152024_2035_vlada-20220511.pdf [8] t. pavlů, j. pešta, m. volf, a. lupíšek. catalogue of construction products with recycled content from construction and demolition waste. iop conference series: earth and environmental science 290(1):012025, 2019. https://doi.org/10.1088/1755-1315/290/1/012025 [9] katalog výrobků a materiálů s obsahem druhotných surovin z průmyslových provozů a komunálních odpadů pro použití ve stavebnictví (2021) – recyklujme stavby!, 2022. [2022-03-03]. http://www.recyklujmestavby.cz/ download/katalog-ii-2021/ [10] produkce, využití a odstranění odpadu a produkce druhotných surovin v roce 2018. [2022-03-03]. https://www.czso.cz/documents/10180/98121450/ 280029-19.pdf/a732554f-bc8f-431a-b41bacd9b89b1f8a?version=1.0 [11] analýza současného stavu vybraných komodit druhotných surovin a jejich zdrojů včetně vize rozvoje daného odvětví, 2018. [2022-03-03]. http://www.asvep.cz/prilohy/51_analyza_ podkladovy-material-pro-pds-cr_1_optimize.pdf [12] statistická ročenka životního prostředí čr 2020, cenia, 2022. [2022-03-03]. https://www.cenia.cz/publikace/statistickarocenka-zivotniho-prostredi-cr/statistickarocenka-zivotniho-prostredi-cr-2020/ 655 https://www.mzp.cz/c1257458002f0dc7/cz/poh_cr_prislusne_dokumenty/%24file/oodp-poh_cr_2015-2024_2035_vlada-20220511.pdf https://www.mzp.cz/c1257458002f0dc7/cz/poh_cr_prislusne_dokumenty/%24file/oodp-poh_cr_2015-2024_2035_vlada-20220511.pdf https://www.mzp.cz/c1257458002f0dc7/cz/poh_cr_prislusne_dokumenty/%24file/oodp-poh_cr_2015-2024_2035_vlada-20220511.pdf https://doi.org/10.1088/1755-1315/290/1/012025 http://www.recyklujmestavby.cz/download/katalog-ii-2021/ http://www.recyklujmestavby.cz/download/katalog-ii-2021/ https://www.czso.cz/documents/10180/98121450/280029-19.pdf/a732554f-bc8f-431a-b41b-acd9b89b1f8a?version=1.0 https://www.czso.cz/documents/10180/98121450/280029-19.pdf/a732554f-bc8f-431a-b41b-acd9b89b1f8a?version=1.0 https://www.czso.cz/documents/10180/98121450/280029-19.pdf/a732554f-bc8f-431a-b41b-acd9b89b1f8a?version=1.0 http://www.asvep.cz/prilohy/51_analyza_podkladovy-material-pro-pds-cr_1_optimize.pdf http://www.asvep.cz/prilohy/51_analyza_podkladovy-material-pro-pds-cr_1_optimize.pdf https://www.cenia.cz/publikace/statisticka-rocenka-zivotniho-prostredi-cr/statisticka-rocenka-zivotniho-prostredi-cr-2020/ https://www.cenia.cz/publikace/statisticka-rocenka-zivotniho-prostredi-cr/statisticka-rocenka-zivotniho-prostredi-cr-2020/ https://www.cenia.cz/publikace/statisticka-rocenka-zivotniho-prostredi-cr/statisticka-rocenka-zivotniho-prostredi-cr-2020/ acta polytechnica ctu proceedings 38:649–655, 2022 1 introduction 2 methods 3 results 3.1 production of secondary raw materials 3.2 possible use of secondary raw materials 3.2.1 slag 3.2.2 ashes and fly ash 3.2.3 energy gypsum 3.2.4 plastic materials 3.2.5 glass 3.2.6 paper 3.2.7 composite packaging 3.2.8 tires and rubbers 3.2.9 wood 3.2.10 metals 4 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0011 acta polytechnica ctu proceedings 34:11–14, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague influence of oxygen and argon plasma treatment on wettability and surface morphology of polypropylene microfibers jakub ďurejea, ∗, zdeněk prošeka, jan trejbala, štěpán potockýb, pavel tesáreka a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b academy of sciences of the czech republic, institute of physics, cukrovarnická 10, 162 53 prague 6, czech republic ∗ corresponding author: jakub.dureje@fsv.cvut.cz abstract. the surface treatment of polypropylene microfibers by plasma processing in oxygen and argon was studied with the motivation to increase their cohesion with the cement matrix. by plasma treatment, it is possible to modify the surface of the microfibers mechanically, chemically or with a combination of both. the microfibers were modified using the reactive ion etching plasma system. the surface of the microfibers was modified by oxygen and argon plasma. the wettability of the microfibers was measured using the packed cell method. furthermore, the weight of the microfibers before and after plasma modification was measured. finally, the surface of the modified microfibers was examined by scanning electron microscopy. almost all modifications of the microfiber surfaces were able to increase their wettability with water. the wettability of the oxygen plasma treated microfibers increased on average by about 11%, the wettability of the argon plasma treated microfibers increased on average by about 6% compared to the fibers without modification. the mechanical effect of plasma treatment was proven only for microfibers modified by oxygen plasma. keywords: oxygen plasma, argon plasma, plasma treatment, plasma modification of microfibers, polypropylene, sem, wettability. 1. introduction the interfacial transition zone significantly influences the properties of cement composite materials containing fibrous reinforcement at the interface between the fiber and the cement matrix [1, 2]. to achieve suitable mechanical properties of the fibers reinforced cement composite material, it is desirable to increase the wettability by water and mainly mechanically modify the fiber’s surface [3]. by plasma modification, it is possible to modify the surface of the fibers both chemically and mechanically. previous experiments proved that plasma modification by oxygen plasma could sufficiently increase the wettability of polypropylene (pp) fibers with water [4, 5]. however, the mechanical effect of plasma modification, realized by inductively coupled plasma, was insufficient on the surface of microfibers [4, 5]. the increased physical effect of plasma modification can be achieved by the impact of energetic ions i.e., with plasma system generally used for ion etching which utilizes ion bombardment process [6]. a typical representative of such systems is the reactive ion etching (rie) system intensively etching the material’s surface [7]. to increase the physical modification of pp microfibers, we present a study of oxygen and argon pp microfibers modifications by the rie system to their wettability and surface morphology. 2. materials and methods polypropylene microfibers fibrifor multi (contec fiber ag, switzerland) were used. the microfibers are made in bundles (type 127) with the diameter of each individual microfiber around 32 µm and the length of 12 mm ± 5 %. the properties of the microfibers are summarized in table 1. the surface of the fibers was modified using the rie system (phantom iii, trion technology inc., usa). microfibers weighing approximately 5 g were inserted into a petri dish. three petri dishes with microfibers were inserted into the rie system and pumped down to the pressure <1·10−3 pa. after that, the chamber was filled (50 sccm) by selected gas (oxygen or argon) and maintained at pressure 13.3 pa. the plasma modification time was set from 30 to 480 seconds. the influence of rf power on microfibers was tested in the range of 100 w to 500 w for a fixed treatment time of 60 seconds. the influence of working gas atmosphere was studied at fixed rf power 100 w and various treatment times were studied. after the plasma modification process, the chamber was evacuated, vented with nitrogen and the microfibers were removed. the microfibers were stored in airtight bags. to determine the treatment stability, part of the samples was exposed to ambient conditions in a laboratory environment at a temperature of about 22 ◦ c and relative 11 https://doi.org/10.14311/app.2022.34.0011 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en j. ďureje, z. prošek, j. trejbal et al. acta polytechnica ctu proceedings humidity of about 50 % for 12 hours. the sample description with relevant process parameters of the rie is in table 2. fibers fibrofor multi material polypropylene diameter [µm] 32 length [mm] 12 tensile strength [mpa] 270 melting point [°c] 160 bulk density [g/cm3] 0.91 table 1. properties of selected fibers. designation input power [w] modification time [s] gas a200w 200 60 argon a300w 300 60 argon a400w 400 60 argon a500w 500 60 argon a30 100 30 argon a60 100 60 argon a120 100 120 argon a240 100 240 argon a480 100 480 argon o30 100 30 oxygen o60 100 60 oxygen o120 100 120 oxygen o240 100 240 oxygen o480 100 480 oxygen table 2. designation and device settings during plasma treatment of microfibers. all microfibers with the petri dishes were weighed on a laboratory scale denver instrument s-64 (denver instrument company, usa) before and after plasma modification. the wettability of microfibers was measured by the packed-cell method. the microfibers were in a perforated container and loaded by a lead weighing 53.55 g. the weight of the container was 9.45 g. subsequently, the sample (microfibers, container, and lead) was weighed on a kern abs (kern & sohn gmbh, germany) laboratory scale. weight values were rounded to tens of mg. next, the sample was dipped into the water for 10 seconds, removed from the water and placed in the holder for 120 s, during which time excess water could drain from the container. finally, the sample was weighed again. from the measured weights, the percentage weight of water compared to the weight of microfibers was calculated according to equation 1 [8]. mv = (mm − mn) − (ms − mn) (ms − mn) · 100, (1) where: mv – weight of water to weight of microfibers [%], mm – weight of wet microfibers with container and lead [g], ms – weight of dry microfibers with container and lead [g], mn – weight of a wet container with wet lead [g]. the surface morphology of the microfibers was examined by electron scanning microscopy (sem, phenom xl, usa). to observe objects using sem, the surface of these objects must be conductive. for this reason, a thin layer of platinum was applied to the surface of the microfibers using a mini sputter coater quorum sc7620 (quorum technologies, uk). platinum was sputtered at a pressure of 8 pa for 120 seconds resulting in approx. 10 nm thin pt layer. the sem was set to high resolution (10 kv) with low current intensity. a sed detector was used. the high vacuum option on the sem was chosen, this option set pressure in the chamber of the device to 1 pa. images of fiber surfaces from sem were magnified 10,000×. 3. results and discussion the surface of the fibers was modified by oxygen plasma and by argon plasma. a more pronounced effect of mechanical modification was expected with argon plasma treatment because argon ions have a higher mass than oxygen ions [9]. in the preliminary runs we tested different rf power inputs from 100 w to 500 w at constant time of 60 s in argon atmosphere. as the power input increases, the temperature loading also increases. at a power input of 300 w and higers (samples a300w, a400w and a500w), the microfibers fused together by heat and started to melt. the temperature of the microfibers during the modification has to reach their melting point, which is around 160 ◦c. due to this fact further testing was performed at a fixed rf power of 100 w. at a fixed rf power of 100 w, we tested different times from 30 s to 480 s in argon or oxygen atmosphere. the weight loss of the microfibers during the plasma modification process was almost negligible, the maximum weight loss was about 0.12 % (sample o480). for microfibers modified by oxygen plasma for 240 s and 480 s, the weight loss was significantly higher compared to other types of plasma modification. weight loss is mainly caused by etching of the microfiber surface during modification. therefore, we can conclude that oxygen plasma changes the morphology of the microfiber surface more than argon plasma. the summary of microfibers weight losses after the plasma modification is shown in figure 1. the wettability of the microfibers was slightly higher after almost all performed plasma modifications. the weight of water between the microfibers in the samples after argon plasma treatment increased on average by about 6 % compared to the microfibers without modification. the weight of water between 12 vol. 34/2022 influence of oxygen and argon plasma treatment the microfibers in the samples after oxygen plasma treatment increased on average by about 11 % compared to the microfibers without modification. the microfibers that were left in the laboratory environment for 12 hours (12h added to sample description) after the plasma modification had similar results of wettability as measured immediately after the plasma modification. the weight of water between the microfibers to weight of microfibers is given in figure 2. morphological changes were observed on the surface of microfibers modified by oxygen plasma for 60 seconds or more by the sem. with increasing time, the mechanical effect of plasma modification due to the ion bombardment was increased. in the sem images we can observe that the mechanical modifications occurred preferentially in one direction. the most pronounced changes were observed on top of the samples (microfibers) placed in the chamber of the rie system. the typical sem images are shown in fig. 3. no morphological changes were observed by sem on the surfaces of the argon modified fibers, so not only the ion bombardment process but also the chemical effect of plasma modification is important to achieve morphological changes on the surface of microfibers. according to the results, to achieve more significant morphological changes, it may be appropriate to extend the modification time, increase the rf power of the rie system or use a mixture of gases (oxygen and argon) during the modification. figure 1. percentage weight loss of microfibers after argon and oxygen plasma treatment (a, o) and respective time in seconds (30, 60, 120, 240 and 480). 4. conclusion plasma modifications of polypropylene microfibers was performed using the rie system and compared the wettability and mass changes depending on plasma process parameters. we studied the influence of, plasma modification times (30–480 s), and working gases (argon or oxygen). based on the performed measurements and results, we can conclude: • polypropylene microfibers begin to melt when the power input of the plasma modification device is set to 300w or more for ar plasma and tested time of 60 s. figure 2. wettability measured by the indirect packed cell method – reference microfibers (ref) microfibers after argon and oxygen plasma treatment (a, o), respective time in seconds (30, 60, 120, 240 and 480) and samples that were tested after 12 hours in a laboratory environment (12h). figure 3. sem images of treated microfibers: a) reference (ref), b) o60, c) o120, d) o240, e) o480. • plasma modifies only the microfiber surfaces. the highest weight loss did not exceed 0.12 % of the initial weight, so that the cross-sectional area of the microfiber almost did not decrease. • at a rf power of 100 w, modification with oxygen plasma for 240 s and 480 s is most suitable – the greatest weight losses occurred during the 13 j. ďureje, z. prošek, j. trejbal et al. acta polytechnica ctu proceedings modification, while the reduction of the microfiber cross-sectional area is negligible. • a slight increase in wettability occurred in almost all samples after plasma treatment. there was on average a more significant increase in wettability of the microfibers comparing oxygen to argon plasma modification. leaving the microfibers in a laboratory environment for 12 hours after plasma modification had no significant effect on their wettability. • significant mechanical changes were observed on the surface of the oxygen plasma treated microfibers at 10,000× magnification. with increasing plasma modification time, the changes were more pronounced. the first mechanical changes have already been observed on fibers treated at 100 w after 60 s. • no morphological changes were observed on the surface of the argon plasma treated microfibers at 10,000× magnification. treatment of the microfiber surface only by argon plasma appears to be inappropriate. • the morphological modification of the microfibers by oxygen plasma was very different on each section of the fiber surface, the fibers were mostly modified only from one direction. to achieve a uniform surface treatment, it would be necessary to carry out the plasma modification several times and to mix the fibers between each modification or to use a mixer during the plasma processing. the results show that the mechanical effect of the plasma modification depends not only on the mass of ions, but also significantly depends on the chemical processes that occur during the plasma modification. maybe it could be achieved better results by using a mixture of oxygen and argon gas during the plasma modification process. better results could be also achieved by increasing the rf power or extending the modification time. 5. acknowledgements this work was financially supported by the czech technical university in prague – the project sgs19/148/ohk1/3t/11 and tačr ss03010302. the authors also thank the center for nanotechnology in civil engineering at the faculty of civil engineering, czech technical university in prague. references [1] k. scrivener, a. crumbie, p. laugesen. the interfacial transition zone (itz) between cement paste and aggregate in concrete. interface science (12):411–421, 2004. https://doi.org/10.1023/b:ints.0000042339.92990.4c. [2] v. machovic. microstructure of interfacial transition zone between pet fibres and cement paste. acta geodynamica et geomaterialia pp. 121–127. https://doi.org/10.13168/agg.2013.0012. [3] a. bentur, m. sidney. fibre reinforced cementitious composites. new york: taylor, 2007. [4] r. hlůžek, z. prošek, j. trejbal, et al. plasma treatment impact on physical and chemical properties of polymeric fibers. acta polytechnica ctu proceedings 13:49–54, 2017. https://doi.org/10.14311/app.2017.13.0049. [5] j. ďureje, z. prošek, j. trejbal, et al. plasma treatment impact on physical and chemical properties of polymeric fibers. acta polytechnica ctu proceedings 13:49–54, 2021. https://doi.org/10.14311/app.2021.30.0007. [6] m. collaud, s. nowak, o. küttel, et al. surface modifications of polypropylene after in-situ ar and n2 plasma treatments: an xps study. applied surface science 72(1):19–29, 1993. https://doi.org/10.1016/0169-4332(93)90038-d. [7] triontech. phoenix. https://triontech.com/. [8] j. ďureje. recycling and its use in concrete waste processing by high-speed milling, 2018. diplomathesis. [9] p. slepička, a. vasina, z. kolaská, et al. argon plasma irradiation of polypropylene. nuclear instruments and methods in physics research 268(11–12):2111–2114, 2010. https://doi.org/10.1016/j.nimb.2010.02.012. 14 https://doi.org/10.1023/b:ints.0000042339.92990.4c https://doi.org/10.13168/agg.2013.0012 https://doi.org/10.14311/app.2017.13.0049 https://doi.org/10.14311/app.2021.30.0007 https://doi.org/10.1016/0169-4332(93)90038-d https://triontech.com/ diploma thesis https://doi.org/10.1016/j.nimb.2010.02.012 acta polytechnica ctu proceedings 34:11–14, 2022 1 introduction 2 materials and methods 3 results and discussion 4 conclusion 5 acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0093 acta polytechnica ctu proceedings 40:93–97, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague application of cold atmospheric plasma for mold inactivation petra ticháa, ∗, mária domonkosa, zuzana rácováb, pavel demoa a czech technical university in prague, faculty of civil engineering, department of physics, thákurova 7, 166 29 prague 6, czech republic b czech technical university in prague, faculty of civil engineering, department of architectural engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: petra.ticha@fsv.cvut.cz abstract. mold growth in indoor interiors is an increasing problem that has adverse effects on both occupants and building materials. the aim of this study was to present the new environmentally friendly method (cold atmospheric plasma treatment) for decontamination of indoor environments contaminated by mold spores, with the potential to become an effective replacement for conventional methods, such as ultraviolet-c (uv-c) germicidal irradiation. the spores of aspergillus species on malt extract agar plates were exposed to plasma produced by diffuse coplanar surface barrier discharge. the surface coverage of mold was evaluated using image analysis. cold plasma treatment inactivated spores on the surface of the culture medium in 0.5 min, whereas a weak growth delay was observed after 30 minute exposure to uv-c irradiation. this study demonstrated that cold atmospheric plasma is a more effective method for reducing mold spores on agar compared with germicidal uv-c irradiation. keywords: cold atmospheric plasma, mold, spore, inactivation, mold growth curve. 1. introduction people spend most of their time in buildings. the presence of mold contamination in indoor environments has adverse health effects as a result of inhaling mold spores and microbial volatile organic compounds produced from mold metabolism. the most common health risks related to mold exposure are eye, skin, and respiratory tract irritation and also activation of immune responses through the inhalation of airborne mold spores. their concentrated presence in the indoor environment can cause degradation of the building material and changes in the visual appearance of the surfaces [1]. the major factors affecting indoor mold growth are temperature (between 15 and 30 °c), moisture, and nutrients. cladosporium, aspergillus, penicillium, and alternaria are identified as the predominant mold species in indoor environments [2]. one of the effective ways to manage mold in a building is to eliminate or limit mold spores and their growth. for microbiological decontamination/inactivation, many methods (mechanical, chemical and physical) and technologies (e.g., uvgi technology, hepa filtration technology) are used. they are mostly based on the application of chemical substances that are often toxic [3]. furthermore, some spores are highly resistant to physical and chemical agents. in recent years, cold (non-thermal) plasma has emerged as effective sporicide against a wide spectrum of bacterial and fungal spores [4, 5]. cold atmospheric plasma (cap) is widely used in various applications such as surface modification, inactivation of pathogens, degradation of toxins, water purification, etc. [6, 7]. cold atmospheric plasma is known for its decontamination and disinfection effects against bacteria, viruses and fungi [4]. several studies have reported that cap can effectively inactivate microorganisms in the food industry, medicine (e.g., prevention of food spoilage, foodborne diseases, sterilization of medical devices) [8]. cap also seems to represent a possible alternative method to standard techniques for the inactivation of mold spores, bacteria, viruses, mycotoxins and fungi in the indoor environment or the mold growth on the surfaces of building materials. cap is an ionized gas that contains a wide range of energetic species, such as reactive oxygen and nitrogen species (rons), free electrons, neutral molecules, and ultraviolet radiation. cold atmospheric plasma can be generated from air at atmospheric pressure. atmospheric plasma equipment is significantly cheaper compared to low pressure plasma system because it does not require expensive vacuum chambers and pumps. plasma technology does not use toxic chemicals. these make plasma treatment very fast, economically advantageous and environmentally friendly [8]. the discharge methods for generating cap mainly include, plasma jets, dielectric barrier discharge, floating-electrode dielectric barrier discharge and corona discharge [8]. the effectiveness of cap in inactivating spores/microorganisms is affected by various factors such as plasma reactor type, device setup, gas composition, operating conditions (e.g., input power) and treatment duration. ultraviolet-c (uv-c) light is a short wavelength (200–280 nm) radiation and is primarily used for the 93 https://doi.org/10.14311/app.2023.40.0093 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en p. tichá, m. domonkos, z. rácová, p. demo acta polytechnica ctu proceedings figure 1. photo of plasma treatment of aspergillus brasiliensis on malt extract agar plates using diffuse coplanar surface barrier discharge (dcsbd) in ambient air. disinfection of air and surfaces from microbial contaminants. in this study, the efficacy of mold spore inactivation by uv-c irradiation and cold atmospheric plasma is demonstrated. 2. materials and methods 2.1. mold and growth conditions the mold strain used in this study was aspergillus brasiliensis ccm8222 (atcc 16404). a volume of 0.1 ml of spore suspension (26 micrormycetes/0.1 ml) was inoculated onto one surface of each test specimen using a pipette. the aim of pipetting the suspension on the surface was to inoculate the spore suspension in the center of the petri dishes (diameter 90 mm). in an effort to minimize contamination with spores and dust, the inoculation was conducted in a laminar airflow cabinet and the test specimens were manipulated with gloves. after inoculation, the samples were incubated for 14 days under constant conditions at 23 °c and 86.7 ±1.7 % relative humidity. 2.2. mold inactivation the inoculated samples were treated in a diffuse coplanar surface barrier discharge (dcsbd) developed by robust plasma systems (rps400; roplass, czech republic). the dcsbd consists of a large number of microdischarges and generates a plasma layer with approximately 0.3 mm thickness on the surface of the planar electrode. the area of the planar electrode connected to the hv power supply is 80 × 200 mm. the frequency of supply voltage is usually in the range of 14–18 khz and the corresponding discharge input power ranges from 80 w up to 400 w. the specific property of the dcsbd plasma is the high plasma power density of 100 w · cm−3 and the homogeneity of the dcsbd plasma increasing with the discharge power density. the dcsbd plasma source was applied to the surface of samples with spores for an exposure time ranging from 0.5–10 minutes with a power of 300 w, and to the surface of samples with young mycelium (treated 72 hours after inoculation) for an exposure time of 10 min with a power of 300 w. all figure 2. image analysis of aspergillus brasiliensis on malt extract agar plates using fiji software. samples were treated homogeneously. the plasma treatment of the sample is shown in figure 1. ultraviolet (uv-c) irradiation was performed using a germicidal uv-c lamp (lb 301.2, 2 × 30 w, tubes length 920 mm, wavelength 253.7 nm). the intensity was approximately 450 µw/cm2 across the exposure area. all samples in the uv-c irradiation group were exposed for 30 min and the distance between the lamp and the samples was 30 cm. 2.3. mold growth assessment the mold coverage on each of the plates was determined by image analysis using the open-source fiji software. a digital photograph of the samples was taken using a canon g1x camera. an iso speed (camera sensor’s sensitivity to light), a shutter speed, and an aperture were set to 800, 1/100 s, and f /5.6, respectively. all pictures of the petri dishes were taken under the same lighting conditions to avoid reflections from the petri dishes surfaces. a per-image thresholding method was used to detect the mold growth from its background (growth media) and resulted in a more accurate measurement of mold area coverage (figure 2). imagej is a useful tool for evaluating mold growth on malt extract agar plates. image analysis provides an objective evaluation and improves the comparability of results. the accuracy of the image analysis was increased by using the same lighting conditions. mold growth on the inoculated surface of each test sample was evaluated once a day for 14 days. the percentage of mold coverage c was calculated according to equation (1): c = am a · 100%, (1) where am is the area of the plate surface that was covered by mold, a is the total area of the petri dish. 94 vol. 40/2023 application of cap for mold inactivation figure 3. mold growth of aspergillus brasiliensis on malt extract agar plates. representative pictures were taken 3, 5, 7 and 9 days after inoculation: (a) control sample, (b) uv-c treated sample (exposure time 30 minutes, treated 24 hours after inoculation), (c) plasma treated sample (300 w, exposure time 0.5, 1, 5 or 10 minutes, treated 24 hours after inoculation), (d) plasma treated sample (300 w, exposure time 10 minutes, treated 72 hours after inoculation). 3. results and discussion using aspergillus brasiliensis spores as a target organism, a comparison of low temperature-based decontamination techniques is reported. the inactivation of a. brasiliensis was performed using cap and uv-c irradiation and was determined by image analysis of the mold growth area in petri dishes. there were six treatment groups and one control group (figure 3). in the case of control group, a. brasiliensis grew rapidly after inoculation and the mold completely covered the petri dishes after 12 days. the mold growth was noticeable 2 days after inoculation, the molds grew from the center to the outer edge and formed a circular colony with irregular edges (mold spores were inoculated at the center of the nutrient). the control sample (figure 3) shows the mold growth phases. initial spore inoculation is followed by lag phase (spore adhesion), spore germination, hyphae growth and reproduction. during the lag phase, the spores remain inactive until they have absorbed sufficient moisture and nutrients from the substrate. hyphae are filaments that appear immediately after germination, and as they thicken, they form a mass called mycelium. colonies of a. brasiliensis were initially yellowish and then dark brown to black. in figure 3, morphological differences were observed between the control and cap treated samples, such as color changes and colony growth arrest. cold plasma treatment of the samples (72 hours after inoculation) for 10 min with a power of 300 w completely inhibited spore germination and hyphal growth of a. brasiliensis. as demonstrated in figure 3, 30-minute uv-c treatment, slightly prevented the colonization of a. brasiliensis on samples. in inactivation spore experiments (24 hours after inoculation), plasma treatments were performed for 0.5, 1, 5 and 10 minutes at 300 w. plasma treated samples did not show mold growth during 14 days. the biological effects of plasma treatment have been demonstrated in several studies [4, 8, 9]. several important mechanisms occur in the process of plasma inactivation, such as chemical reactions with atomic oxygen, and uv-induced damage. the microbial inactivation of plasma is mainly attributed to rons. rons include o3 and h2o2, which are powerful oxidants and have similar inactivation mechanisms. these reactive species inactivate molds by progressive oxidative damage to their cell walls and membranes. uv radiation, which is also produced during the formation of cap, is not considered to be a dominant factor in the inactivation process. the results of mold growth image analysis during 14 days of incubation are depicted in figure 4. all data are presented as the mean ± standard deviation. 95 p. tichá, m. domonkos, z. rácová, p. demo acta polytechnica ctu proceedings figure 4. mold growth curves of aspergillus brasiliensis on malt extract agar plates. the mold in the control samples grew according to the typical s-shaped curve (logistic growth). the lag phase lasted two days. thus, the measurements started on the second day and continued until the 14th day. the exponential growth went up to the 7th day (inflection point) and it was possible to observe a color change, to a black tone in the center of colonies. compared to the control group, mold growth was negligibly suppressed by the uv-c germicidal lamp. as shown in figure 4, the coverage area of aspergillus brasiliensis treated with uv-c irradiation was slightly smaller than that of the control group at the same time. spore color can play an important role in inactivation of mold spores by uv-c irradiation. the aspergillus brasiliensis spores are more resistant, in that the pigment of the a. brasiliensis spores absorbs more in the uv-c region and thus protects the spores from uv irradiation [10]. the growth curve results also clearly indicate that the 30-second cap treatment (24 hours after inoculation) was sufficient to completely inactivate the a. brasiliensis spores and prevent mold growth in all samples. the effect of cap treatment on permanent inactivation of a. brasiliensis spores was also evaluated for longer exposure times (1–10 min). based on the above results, plasma treatment is considerably more effective than a conventional uv-c irradiation. cap treatment achieved a high level of inactivation level within a short time. 4. conclusion in this study, it can be concluded that the cold atmospheric pressure plasma has potential sporicidal activity against molds and could be a promising tool for the inactivation of spores in indoor environments or on the surface of building materials. the cold atmospheric plasma treatment was compared with a conventional germicidal uv-c lamp, which is commonly used to reduce mold spores in indoor environments. the results of image analysis of mold growth on malt extract agar plates treated with germicidal uv-c lamp showed a non-significant delay in mold growth dynamics. the appearance of the first mold colonies exposed to uv-c irradiation was delayed by only 24 hours compared to the control samples. furthermore, 0.5 min of cap treatment (24 h after inoculation) was shown to lead to inactivation of a. brasiliensis spores. the inactivation of young mycelium (hyphae) was also tested by cold atmospheric pressure plasma treatment. images of samples with young mycelium exposed to plasma produced by diffuse coplanar surface barrier discharge for 10 minutes at 300 w (72 hours after inoculation) show an obvious change in mold color and an arrest of growth. compared to control samples, it was observed that cold atmospheric pressure plasma treatment caused inactivation of spores and young mycelium growth. the mode of cap action on microorganisms is still the subject of extensive studies, but it is believed that biologically reactive substances generated by plasma (rons, uv radiation, and charged particles) act cooperatively to inactivate microorganisms. however, more studies are needed to understand in detail the mechanism of spore inactivation and young mycelium growth by cold atmospheric plasma. 96 vol. 40/2023 application of cap for mold inactivation acknowledgements this work was supported by the czech science foundation (no. ga22-06621s). references [1] p. ryparová, r. wasserbauer, z. rácová. the cause of occurrence of microorganisms in civil engineering and the dangers associated with their growth. procedia engineering 151:300–305, 2016. https://doi.org/10.1016/j.proeng.2016.07.402 [2] c. du, b. li, w. yu. indoor mould exposure: characteristics, influences and corresponding associations with built environment—a review. journal of building engineering 35:101983, 2021. https://doi.org/10.1016/j.jobe.2020.101983 [3] a. m. elkhatat. recent trends of microbial decontamination for occupational, industrial and domestic applications. bulletin of the national research centre 46(1):75, 2022. https://doi.org/10.1186/s42269-022-00761-9 [4] v. scholtz, j. pazlarová, h. soušková, et al. nonthermal plasma — a tool for decontamination and disinfection. biotechnology advances 33(6):1108–1119, 2015. https: //doi.org/10.1016/j.biotechadv.2015.01.002 [5] b. g. dasan. decontamination of aspergillus f lavus and aspergillus parasiticus spores on hazelnuts via atmospheric pressure fluidized bed plasma reactor. international journal of food microbiology 216:50–59, 2022. https: //doi.org/10.1016/j.ijfoodmicro.2015.09.006 [6] m. domonkos, p. tichá. wettability of siox nanofibrous mats and their modification using cold atmospheric plasma treatment. acta physica polonica a 140(2):181–185, 2021. https://doi.org/10.12693/aphyspola.140.181 [7] j. trejbal, v. nežerka, m. somr, et al. deterioration of bonding capacity of plasma-treated polymer fiber reinforcement. cement and concrete composites 89:205–215, 2018. https: //doi.org/10.1016/j.cemconcomp.2018.03.010 [8] m. domonkos, p. tichá, j. trejbal, p. demo. applications of cold atmospheric pressure plasma technology in medicine, agriculture and food industry. applied sciences 11(11):4809, 2021. https://doi.org/10.3390/app11114809 [9] e. lokajová, j. julák, j. khun, et al. inactivation of dermatophytes causing onychomycosis using non-thermal plasma as a prerequisite for therapy. journal of fungi 7(9):715, 2021. https://doi.org/10.3390/jof7090715 [10] v. m. gómez-lópez, p. ragaert, j. debevere, f. devlieghere. pulsed light for food decontamination: a review. cement and concrete composites 18(9):464–473, 2007. https://doi.org/10.1016/j.tifs.2007.03.010 97 https://doi.org/10.1016/j.proeng.2016.07.402 https://doi.org/10.1016/j.jobe.2020.101983 https://doi.org/10.1186/s42269-022-00761-9 https://doi.org/10.1016/j.biotechadv.2015.01.002 https://doi.org/10.1016/j.biotechadv.2015.01.002 https://doi.org/10.1016/j.ijfoodmicro.2015.09.006 https://doi.org/10.1016/j.ijfoodmicro.2015.09.006 https://doi.org/10.12693/aphyspola.140.181 https://doi.org/10.1016/j.cemconcomp.2018.03.010 https://doi.org/10.1016/j.cemconcomp.2018.03.010 https://doi.org/10.3390/app11114809 https://doi.org/10.3390/jof7090715 https://doi.org/10.1016/j.tifs.2007.03.010 acta polytechnica ctu proceedings 40:93–97, 2023 1 introduction 2 materials and methods 2.1 mold and growth conditions 2.2 mold inactivation 2.3 mold growth assessment 3 results and discussion 4 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0310 acta polytechnica ctu proceedings 38:310–315, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague experimental monitoring of autonomous curtain walling facade module vojtěch zavřela, b, ∗, tomáš matuškab, petr slaninac a czech technical university in prague, faculty of mechanical engineering, department of environmental engineering, technická 4, 166 07 prague 6, czech republic b czech technical university in prague, university centre for energy efficient buildings, department of energy systems in buildings, třinecká 1024, 273 43 buštěhrad, czech republic c wieden s.r.o., do čertous 2830/2, 193 00 prague, czech republic ∗ corresponding author: vojtech.zavrel@fs.cvut.cz abstract. this paper introduces an innovative concept for active wall-curtain façade modules aiming at high level of energy autonomy. the façade module consists of two opaque panels and one transparent panel integrating building integrated photovoltaics supported by façade integrated battery peltier air-conditioning unit, active shading, and led lighting. the developed demo-protype was deployed at testing facility and subjected to long-term experimental monitoring to assess the energy performance in real environment. the complex monitoring system captured key power and thermal fluxes, temperatures, mass flows and other operational states of the active wall-curtain façade modules. these measurements were assessed in form of monthly overview depicting total electricity consumption, on-site energy production, heating and cooling delivery as well as self-sufficiency and self-consumption indicators. keywords: active facades, energy autonomy, bipv, in-façade energy systems, peltier air-conditioning. 1. introduction the eu committed in the european green deal to decarbonize the building sector by 2050 [1]. however, according to bpie [2], over 97 % of buildings around eu have to undergo the retrofit to achieve this vision. they also stated that 75 % of the building stock are built before 1990 in outdated envelope standards leading to insufficient energy performance. this situation calls for high-performance façade solutions enabling to achieve nzeb standard for new or retrofitted buildings. the building envelope must cope with daily and seasonal variation in both interior and exterior environment due to the occupancy requirements and given weather conditions. however typical envelope has usually static properties, with lack of flexibility to reflect these variations or harvest the incident solar radiation. this paper introduces compact façade solution embedding multiple energy systems within the wallcurtain façade module aiming at high level of energy autonomy. the current research proposes an innovative active façade solution integrating bipv panels, flat battery, peltier air-conditioning (ac) unit, active shading system and artificial lighting embedded into one direct current (dc) microsystem developed by czech technical university in prague. this dc microsystem has been integrated into prefab curtainwalling module manufactured by wieden s.r.o., the industry partner within the project. the active façade solution may improve the local energy management, where the solar energy is locally utilized for the space-conditioning or other integrated energy systems. the current work represents experimental research, where the research institution and industry partner closely cooperated in the development of a demo-prototype of the active multi-functional façade curtain walling unit. the demo-prototype was deployed on testing facility located at the uceeb, ctu in prague (see in figure 1) and energy behaviour of the façade unit is being monitored and tested in real environment. so far, the current version of the demo-prototype does not represent a marketable product. the presented façade unit is developed mainly for demonstration purposes of the compact in-façade solutions and embedded energy system cooperation. the development is still in progress to reach the marketable level, that the active curtain walling module, complies all required standards. the demonstration façade unit is compounded from three main modules: an opaque module with bipv and integrated flat-plate batteries, an opaque module with bipv and façade integrated peltier ac unit and a transparent module with bipv and active shading device. the technical specification of these panels is provided in sections below. hereunder, the applied technologies are briefly discussed. the bipv technology was utilized as a main source of energy to supply the active façade unit. this technology represents the best-practice building component, that has been already established at the market as a promising solution for building stock decarbonization. for further information, this technology has been already comprehensively reviewed in several publica310 https://doi.org/10.14311/app.2022.38.0310 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 experimental monitoring of autonomous curtain . . . figure 1. demonstration façade curtain walling unit deployed at testing facility. tion e.g. recently in [3, 4]. the intermittent operation of the bipv panels is in our concept supported by (i) prioritizing the direct use by the other energy systems within the façade unit and then (ii) storing the surplus energy in the flat-plate battery array integrated in the inner wall of the unit. the battery integration into the facade has not been as heavily researched as bipv. there are only limited literature sources available regarding such integration. e.g. kim et al. [5] reported integration of a solar power bank for user appliances. the reason is likely due to several practical issues related with questionable fire safety as well as high sensitivity of battery performance for operational conditions. these practical issues were neglected in this research in order to arrive to a compact solution demonstrating the potential of the energy autonomous concept. the next integrated system is façade integrated air-conditioning system. the design of façade unit assumes the integration of air-conditioning device into the curtain walling structure despite the fact, that the potential space for the unit is very limited. the conventional compressor systems cannot fit into the façade structure easily and vibro-acoustic properties together with limited possibilities for maintenance of mechanical components finally exclude such solution from design. peltier cells compiled from thermocouples can be regarded as promising alternative for such application. peltier cells with use of electricity absorb heat on cold side of the thermocouple and reject the heat on hot side [6]. such non-mechanical heat pump offers number of advantages such as minimal dimensions, no mechanical parts, no maintenance, no risk of leakage, suitability for direct current supply from pv system and simple control (heating and cooling mode based on electric current polarity). compared to compressor system, lower efficiency and a limited temperature drop is referenced as main disadvantages [7, 8]. peltier technology utilization in building technical systems has been researched in last decade and number of prototypes has been developed. the first peltier system for heating and cooling purpose integrated into building façade has been developed in 2016 [7]. the combination of a façade-integrated peltier airconditioning device with pv system has been reported a year later [9]. 2. autonomous façade curtain walling unit design 2.1. module 1: building-integrated pv with flat battery the section view of module 1 is depicted in figure 2a. the bipv panels are integrated into a standard design of curtain-walling unit structure. the thickness of the curtain walling unit together with ventilated cavity is about 400 mm. the allocated place for the flat battery is within the 25 mm thin cavity at the inner side. pv system consists of 8 polycrystalline panels considering 17 % efficiency with total peak power about 920 wp. the bipv operation is supported by flat battery array accounting 2 sections, each with 8 lifepo4 battery cells. the nominal voltage of each cell is at 3.2 v and capacity 60 ah. in total, the array is designed for maximum voltage 25.6 v and storage capacity of 3.1 kwh. in addition, the led strip is located at the inner side. 311 v. zavřel, t. matuška, p. slanina acta polytechnica ctu proceedings (a). module 1 with integrated flatplate battery. (b). module 2 with glazing and active shading. (c). module 3 with façade integrated peltier ac unit. figure 2. unit design – section view of modules. 2.2. module 2: glazing with external shading system the section view of module 2 is depicted in figure 2b. this module represents transparent structure for natural daylight combined with mechanical shading system in the front of the large-format triple glazing. servomotor for external louvers is powered by dc current. the shading system emplacement is covered by the pv panel. the shading system is controlled automatically by the embedded smart controller with respect to outdoor environment to prevent the high cooling load in summer season or harvest the solar gain in winter. 2.3. module 3: façade integrated peltier air-conditioning unit the section view of module 3 is depicted in figure 2c. this module is specific by the façade integrated air conditioning unit using peltier cells. the spaceconditioning is provided by 20 pcs of high-performance peltier cells arranged in two rows with design cooling and heating capacity 640 w, 1280 w, respectively. the rows can be operated separately in two stage control manner. the nominal electric power load is assumed at 750 w. the heat is absorbed/removed from cell surfaces by finned heat sinks located at both side of the cells. the heat is then transferred to the air circulating in exterior and interior duct channels. the circulation of the air is provided by two variable speed control fans. the highest airflow is considered at 800 m3/h. the entire ac unit fits into cavity with thickness of 165 mm. voltage polarity of peltier cells defines the operation mode of ac unit (heating, cooling). the air from indoor space enters to the interior channel and cooled down is delivered back to the space. the extracted heat is rejected to the ambient environment through the hot side of peltier cells via the exterior air channel. the operation in winter mode is reversed. 2.4. monitoring system the demo-prototype is equipped with complex monitoring system. this system measures main energy flows within the embedded energy systems of the façade unit, outdoor and indoor climate conditions (temperature and relative humidity of the air, solar irradiance). the thermal output of the peltier ac unit has been measured indirectly based on fan speed vs mass flow mapping and inlet-outlet temperature difference measured directly. the mapping was performed as one-at-time measurements using velocity probing. the outcome of this preparatory experiment was correlation curve between fan speed vs mass flow, that was further used in the energy performance assessment. in addition, the monitoring system as well as logged consumption of secondary heating and cooling systems in the testing facility. the secondary systems serve as back-up energy sources in case of limited heating or cooling capacity of the peltier ac unit. the measured variables are listed in table 1. 3. energy performance assessment the demo-prototype was deployed at testing facility in september 2020. the installation was followed by commissioning period to configure the control and monitoring system. during commissioning period, a faulty behaviour of the on-site power system was found. the battery system was being overloaded every time, when the second stage of the peltier ac unit was triggered. the electrical current required by peltier cells exceeded the nominal level of the battery system. 312 vol. 38/2022 experimental monitoring of autonomous curtain . . . metering definition unit metering definition unit peltier cells consumption kwh peltier ac unit inlet – interior °c fans consumption kwh peltier ac unit inlet – exterior °c back-up heating consumption kwh peltier ac unit outlet – interior °c back-up cooling – fan consumption kwh peltier ac unit inlet – exterior °c monitoring system consumption kwh back-up cooling – inlet water temperature °c led lighting consumption kwh back-up cooling – outlet water temperature °c total electricity consumption kwh back-up cooling – water mass flow °c pv production kwh indoor temperature °c battery supply kwh indoor relative humidity % pv surplus metering via electric heater kwh outdoor temperature °c outdoor relative humidity % solar irradiance on façade exterior w/m2 table 1. list of measured variables. year / month missing data total el. consumption pv system production heating supply and coverage factor cooling supply and coverage factor [%] [kwhel] [kwhel] [kwht] [%] [kwht] [%] 10/2020 3 268 nan 16 59 0.4 17 11/2020 0 322 24 24 71 0.0 n/a 12/2020 0 361 15 145 90 0.0 n/a 01/2021 17 352 21 192 95 0.0 n/a 02/2021 21 273 33 135 74 0.0 n/a 03/2021 2 278 43 136 96 0.8 100 04/2021 0 193 45 73 95 0.4 100 05/2021 0 102 49 13 75 0.0 n/a 06/2021 0 166 51 0 0 4.1 15 07/2021 0 48 42 0 55 0.1 100 08/2021 0 59 34 2 70 0.2 100 09/2021 0 44 25 3 80 0.0 n/a table 2. energy performance of the façade module: monthly overview. thus, the presented results were obtained without using second stage of the ac unit. regarding the operational regime, the heating setpoint of the peltier ac unit was set to 21 °c. the cooling setpoint 27 °c to keep the indoor environment of the testing facility within this range. flowrate through interior and exterior channel of peltier ac unit were different for winter and summer operation. the winter flowrate was approx. 600 m3/h (3800 rpm) for interior channel (fan) and approx. 300 m3/h (1900 rpm) for exterior channel (fan). the summer flowrate was approx. 450 m3/h (2800 rpm) for interior channel (fan) and approx. 650 m3/h (4200 rpm) for exterior channel (fan). in this experiment, secondary energy sources were also in operation. the settings aim to switch on the secondary heat sources only if the peltier ac unit reach its maximum capacity. therefore, the heating setpoint was set at 18.5 °c and cooling setpoint was 28.5 °c. the resulted energy performance is depicted as monthly overview in table 2. the table summarizes the total electrical consumption, measured pv system production and heating/cooling supply and associated coverage factor. remark on share of missing data due to lost internet connection in january and february 2021 is also added. the time-step of the monitoring data acquisition was 30 s and data were logged into cloud archive. detailed energy balance has been monitored for period from october 2020 to september 2021. the pv system electricity production was not measured in october 2020 due to fault of the battery system. various operation parameters have been adjusted within the monitoring and evaluation period. in november 2020, the winter flowrates for channels have been finally set. the summer flowrates for channels wre set in the beginning of may 2021. external mechanical shading was shut in june and the back-up system for cooling was switched off. the level of autonomy and energy efficiency was further processed using following energy performance metric. self-sufficiency ss indicator and selfconsumption indicator sc were used to indicate the level of energy autonomy. some literature denotes these indicators also as on-site energy fraction (oef) and on-site energy matching (oem). indicators ss and sc are defined between 0 to 1, where ss indicates 313 v. zavřel, t. matuška, p. slanina acta polytechnica ctu proceedings year / month missing data ss sc scop max. cop seer max. eer [%] [%] [%] [-] [-] [-] [-] 10/2020 3 n/a n/a 0.08 0.99 0.34 0.39 11/2020 0 8 98 0.10 0.46 0.05 0.17 12/2020 0 4 100 0.46 2.16 n/a n/a 01/2021 17 6 100 0.62 3.32 n/a n/a 02/2021 21 12 1.0 0.57 1.13 n/a n/a 03/2021 2 16 99 0.58 1.75 0.17 0.22 04/2021 0 24 99 0.46 1.91 0.20 0.24 05/2021 0 48 97 0.18 0.64 0.00 0.17 06/2021 0 31 99 0.16 0.38 0.05 0.33 07/2021 0 86 90 n/a n/a 0.01 0.11 08/2021 0 56 91 0.08 0.26 0.01 0.11 09/2021 0 57 92 0.09 0.25 n/a n/a table 3. energy autonomy and efficiency: monthly overview. (a). cop factor and heating load. (b). eer factor and cooling load. figure 3. peltier ac unit efficiency vs outdoor air temperature. the portion of power consumption, that is covered by the on-site source and sc indicates the portion of generated energy from the on-site source that is used within the system, rather than exported or dumped. the peltier ac unit provides the heating and cooling in the testing facility. this device also represents the major electricity consumer of the active facade unit. therefore, this study is focused on evaluation of the energy efficiency using cop for heating and eer for cooling. the monthly summary is shown in table 3. this table contains timestamp, percentage of missing data, self-sufficiency and self-consumption indicators and cop and eer factors. cop and eer factors were evaluated in two forms: (a) as ratio of “seasonal” (monthly) sum of thermal energy delivered/removed by the unit vs its electricity consumption accounted in kwh denoted as scop and seer in table 3 and (b) as ratio of nominal heating/cooling load vs its power load accounted in w. maximal values for given months are also depicted in table 3 denoted as max. cop and max. eer. both forms (a) and (b) include fans power. the efficiency of peltier ac unit was further investigated with respect to outdoor air temperature. figure 3 shows a scatter plot, where each dot represents a measured state for which cop or eer was evaluated related to outdoor temperature at the time of the measurements. in addition, each dot was coloured in scale representing the thermal load of the unit at the time of the measurements to give whole information regarding the energy performance of the peltier ac unit. the maximum cop = 3.2 was reached january 2021. the maximum eer = 0.39 was obtained in october 2020 during the full load. the unit in part-load performed at considerably lower efficiency. the seasonal cop is between 0.46 and 0.62 for winter season and the seasonal eer is between 0.11 and 0.33 for summer season. 4. lessons learned after successful laboratory testing of functional samples, the demo-prototype of the façade curtain walling unit was designed, constructed, and deployed at testing-facility located at uceeb (ctu) site, 314 vol. 38/2022 experimental monitoring of autonomous curtain . . . buštěhrad. the energy performance of the façade unit was experimentally evaluated within one year long monitoring campaign to assess the façade curtain walling unit operation in the real environment. hereunder, the presented results are summarized from perspective of: (i) energy autonomy expressed by power selfsufficiency and self-consumption of the testing facility including local pv, (ii) thermal demand coverage by the façade integrated peltier unit, (iii) efficiency in terms of cop and eer of peltier unit (as main power consumer). (a) in terms of energy autonomy, self-sufficiency during summer season and potential for solar assisted cooling has been evaluated. the highest self-sufficiency indicator was reached in july 2021, power load covered by the on-site energy generation was 87 % due to application of external shading. on the other side, high power self-sufficiency indicator in winter season was low, only around 10 % as indicated already in preliminary simulations. (b) the thermal demand coverage was limited due to second stage disallowance in operation of the peltier ac unit. the missing capacity was especially noticeable in cooling mode without shading. as an example, only about 15 % of cooling load in june 2021 was covered by the peltier system. on the other hand, operation within the transition season or with use of external shading resulted in possibility to cover the demand of testing facility space and it reduced the hours of space overheating. although the total capacity of the peltier ac unit was half of originally assumed, the coverage of heating load was high at approximately 80 % in average during the whole winter season (from october 2020 to march 2021). (c) peltier ac unit performed bellow the expectations. based on laboratory measurements, the cop was expected in range of 1.2 to 2 and the eer in range of 0.6 to 1. the unit in the real-life scenario could not reach the expected energy performance. the cop was found between 0.8 to 0.6 for nominal load and eer around between 0.1 to 0.2. the efficiency is worsened due to part-load operation as well as due to higher temperature difference between ambient and interior air temperature. the energy efficiency could be improved by several measures: (i) enabling the full-capacity, (ii) optimization of current-voltage settings of the peltier, (iii) optimization of fan speed. particularly the constant fan speed setting was found inappropriate. as first step, the variable fan speed control needs to be applied to reflect the regime of the unit and ambient temperature. the design of the optimized fan control is the ongoing work. to conclude, this paper introduced demo-prototype of active façade curtain walling unit, that integrates various energy systems. the demo prototype was successfully deployed at testing facility and long-term monitoring campaign was executed. the detailed monitoring system allowed to assess the level of energy autonomy and energy performance of the innovative façade components. acknowledgements experimental monitoring has been supported by the technology agency of czech republic within the project th03020341 autonomous curtain wall panel. references [1] european commission. establishing the framework for achieving climate neutrality and amending regulation (eu) 2018/1999 (european climate law), brussels, 2020. 25 p. [2] buildings performance institute europe. 97 % of buildings in the eu need to be upgraded, 2017. [2021-10-15]. https://www.bpie.eu/wp-content/uploads/2017/12/ state-of-the-building-stock-briefing_dic6.pdf [3] e. biyik, m. araz, a. hepbasli, et al. a key review of building integrated photovoltaic (bipv) systems. engineering science and technology, an international journal 20(3):833–858, 2017. https://doi.org/10.1016/j.jestch.2017.01.009 [4] a. k. shukla, k. sudhakar, p. baredar. recent advancement in bipv product technologies: a review. energy and buildings 140:188–195, 2017. https://doi.org/10.1016/j.enbuild.2017.02.015 [5] a. a. kim, d. a. reed, y. choe, et al. new building cladding system using independent tilted bipv panels with battery storage capability. sustainability 11(20):5546, 2019. https://doi.org/10.3390/su11205546 [6] s. lineykin, s. ben-yaakov. modeling and analysis of thermoelectric modules. in twentieth annual ieee applied power electronics conference and exposition, 2005. apec 2005, vol. 3, pp. 2019–2023. 2005. https://doi.org/10.1109/apec.2005.1453336 [7] m. ibañez-puy, j. bermejo-busto, c. martín-gómez, et al. thermoelectric cooling heating unit performance under real conditions. applied energy 200:303–314, 2017. https://doi.org/10.1016/j.apenergy.2017.05.020 [8] z. liu, l. zhang, g. gong, et al. review of solar thermoelectric cooling technologies for use in zero energy buildings. energy and buildings 102:207–216, 2015. https://doi.org/10.1016/j.enbuild.2015.05.029 [9] k. irshad, k. habib, f. basrawi, b. b. saha. study of a thermoelectric air duct system assisted by photovoltaic wall for space cooling in tropical climate. energy 119:504–522, 2017. https://doi.org/10.1016/j.energy.2016.10.110 315 https://www.bpie.eu/wp-content/uploads/2017/12/state-of-the-building-stock-briefing_dic6.pdf https://www.bpie.eu/wp-content/uploads/2017/12/state-of-the-building-stock-briefing_dic6.pdf https://doi.org/10.1016/j.jestch.2017.01.009 https://doi.org/10.1016/j.enbuild.2017.02.015 https://doi.org/10.3390/su11205546 https://doi.org/10.1109/apec.2005.1453336 https://doi.org/10.1016/j.apenergy.2017.05.020 https://doi.org/10.1016/j.enbuild.2015.05.029 https://doi.org/10.1016/j.energy.2016.10.110 acta polytechnica ctu proceedings 38:310–315, 2022 1 introduction 2 autonomous façade curtain walling unit design 2.1 module 1: building-integrated pv with flat battery 2.2 module 2: glazing with external shading system 2.3 module 3: façade integrated peltier air-conditioning unit 2.4 monitoring system 3 energy performance assessment 4 lessons learned acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0478 acta polytechnica ctu proceedings 38:478–482, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague durability of latent heat storage systems sylva bantová, milan ostrý∗, karel struhala brno university of technology, faculty of civil engineering, institute of building structures, veveří 95, 602 00 brno, czech republic ∗ corresponding author: ostry.m@fce.vutbr.cz abstract. latent heat storage represents technology with significantly higher energy storage density. the thermal energy storage capacity of building structures and storage units integrated into building services contribute to the energy flexibility of buildings. this paper presents results from laboratory experiments focused on the compatibility of heat storage media represented by phase change materials (pcms) with materials of container. material compatibility of selected pcms with the plastics and metals were tested by a long-term experiment. two organic-based and two inorganic-based phase change materials were selected for tests of compatibility with selected metals (aluminium, copper and brass) and plastics (pp-h, pe-hd, and pvc-u). plastic-pcm compatibility was determined by gravimetric method. for evaluation of metal-pcm compatibility, calculation of corrosion rate was applied. the less mass changes and lower penetration of pcms to the matrix was observed for inorganic-based pcms compared to organic-based pcms. in case of compatibility between metals and pcms, the highest values of corrosion rate were calculated for copper immersed in inorganic-based pcms rubitherm sp25. keywords: phase change materials (pcms), compatibility of materials, metal corrosion, mass changes, container, latent heat storage systems, durability, energy flexibility. 1. introduction the situation on the european energy market in 2021 which is accompanied by rising of energy cost is one of the reasons for the demand of larger energy independence and energy supply safety of energy consumers. when fossil fuels such as natural gas or coal are considered to be replaced by renewable energy sources (res), the sufficient energy storage capacity should be employed. the reason is that energy production from res, such as sun, is not constant in time. if the energy in the form of heat should be stored for short or longer time, then the sensible heat storage or latent heat storage technology should be integrated in energy grid. the amount of stored heat in sensible heat storage technology depends on the weight of storage medium, its thermal capacity and change between initial and final temperature. on the other hand, the latent heat storage represents advanced and much more sophisticated approach to heat storage. materials suitable for latent heat storage are called phase change materials (pcms). latent heat storage technology uses the phase change of storage material and solid-liquid transformation is suitable for building applications. due to the phase change, proper encapsulation of pcms must be designed for all heat storage units with reasonable durability. durability and sustainability of each heat storage unit depends on the compatibility between the material of encapsulation and pcms to avoid undesirable leakages of pcms and their interaction with surrounding environment. this paper presents the results of compatibility tests of pcms selected from organic or inorganic group, and plastics and metals that are suitable materials for their encapsulation. 2. materials and methods 2.1. materials a total of four pcms were selected for testing: salt-hydrate-based rubitherm sp22 and rubitherm sp25 (representatives of the inorganic group) and paraffin-based linpar 17 and linpar 1820 (representatives of the organic group). table 1 shows the selected pcm parameters. as all tested pcms have a mean value of peak temperature around 24 °c, the actual material range values of materials is 22–28 °c, this suggests the suitability of specific pcms for integration into systems inside residential buildings due to similar indoor room temperatures. three common metals and plastics were selected for the experiment as potential encapsulation materials for pcms. the reason for the selection was good market availability, the high value of thermal conductivity of metals and the wide use of plastics in the buildings for distribution systems of rainwater or sewer pipes, for storage tanks, etc. two groups of samples were determined for the experiment: • plastic group representatives: pp-h (polypropylene), pvc-u (polyvinylchloride) and pe-hd (highdensity polyethylene); • metal group representatives: aluminium, copper and brass. table 2 shows the parameters of tested metals. 478 https://doi.org/10.14311/app.2022.38.0478 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 durability of latent heat storage systems type latent heat onset temperature peak temperature [j g−1] [°c] [°c] linpar 17 organic 152 21 22 linpar 1820 organic 141 24 27 rubitherm sp25 inorganic 122 18 28 rubitherm sp22 inorganic 145 14 25 table 1. the parameters of selected pcms. composition density thermal conductivity area [g cm−3] [w m−1k−1] [cm−2] aluminium (en aw1050 h111) al99.5 2.70 230 copper (en cw024a) cu-dhp-99.9 8.96 340 41.80 brass (en cw617n) cuzn40pb2 8.73 114 table 2. the selected parameters of the metal samples. 2.2. methods the assessment of compatibility between selected pcms and the metal/plastic samples is based on the weight changes of the samples with the calculation of the evaluation parameters separately for each material. the test procedure is based on the methodology of the american society for testing and materials test g1-03 [1] and consists of the following parts: • preparation of the samples – visual inspection, cleaning (with toluene for metals), dried, clearly labelled; • weighing of samples – the samples were weighed with the analytical balance; • immersion in pcms – the samples were placed into the testing beakers and were poured with liquid pcms, the test beakers were placed into incubator; • thermal cycling in the incubator – the samples were exposed to repeating temperature cycling ranging from 15 °c to 40 °c, the cycle was divided into 4 parts, each lasting for two hours; • removal from pcms: the samples of one set were withdrawn from the testing beakers at the given time: after 7-28-84 days (for the metals) and every seventh day (for the plastic) of exposure to the pcms, one set containing three pieces of samples; • visual check (the changes of appearance) of samples; • cleaning of the samples after removing; • weighing of the samples; • the selection of one sample from each set for subsequent calculations using the median value – to eliminate extreme values of the weight changes; • the calculation of the parameters (the corrosion rate for metal, the percentage change in mass in case of plastic) and evaluation of the samples. 2.2.1. the method for evaluation of the metal-pcm compatibility the corrosion rate (cr) was calculated for each metal sample in contact with the pcms for a specified time, the procedure and the experimental conditions are given in section 2.2. the cr methodology includes monitoring the mass loss of the metal samples relative to the surface area and the immersion time of the samples, cr is defined by equation (1), described in [2, 3]. cr = ∆m a · (t0 − t) , (1) where cr is the corrosion rate in mg cm−2 year−1, ∆m is a mass loss in mg and it is defined by the equation ∆m = m(t0) − m(t), where m(t0) is the initial weight of the sample before immersion to pcms and m(t) is the final weight after the immersion time, a is the surface area of the sample in cm2 and (t0 − t) is the experimental immersion time of the samples in pcms in years. in this study, the calculated corrosion rates of the samples are compared to the classification described in the guide for corrosion weight loss used in the industry [2, 4]. according to the guide, pcms are classified as compatible with the selected metal or incompatible, or compatible for a specific application. the suitability scales with the limit values and the verbal evaluation of cr are shown in table 3. the overall classification of the metal samples is supplemented by the calculation of the cr value defined according to the methodology described in [5, 6]. the calculation of the cr is based on the american society for testing and materials test g-31 [7] as well, which determines the average cr values. the method is based on the weight changes that are caused by exposure of the samples to a corrosive environment during the experiment. this calculation procedure is not possible to use in case of localized corrosion on the surface of samples such as pitting or intergranular corrosion. 479 s. bantová, m. ostrý, k. struhala acta polytechnica ctu proceedings recommendation (verbal evaluation) cr = 50.0 to 999.0 mg cm−2 year−1 not recommended for service cr = 10.0 to 49.9 mg cm−2 year−1 caution recommended, based on the specific application cr = 0.3 to 9.9 mg cm−2 year−1 recommended for long term service table 3. the guide for corrosion weight loss used in industry (the selected limit values) [2–4]. material immersion linpar 17 rubitherm sp25 period (organic group of pcm) (inorganic group of pcm) ∆m cr cr ∆m cr cr [day] [mg] [mg cm−2 year−1] [mm year−1] [mg] [mg cm−2 year−1] [mm year−1] aluminium 7 0.800 1.049 0.00389 0.800 1.024 0.00379 28 0.700 0.219 0.00081 0.900 0.292 0.00108 84 0.500 0.053 0.00020 1.800 0.194 0.00072 copper 7 1.400 1.807 0.00202 4.400 5.400 0.00603 28 1.700 0.525 0.00059 5.800 1.863 0.00208 84 1.600 0.174 0.00019 9.600 1.004 0.00112 brass 7 1.300 1.685 0.00193 2.900 3.882 0.00445 28 2.000 0.621 0.00071 3.200 1.061 0.00121 84 0.400 0.044 0.00005 4.500 0.476 0.00055 table 4. the corrosion rates values of tested metals exposed to effects of selected pcms from the organic and inorganic group for periods 7, 28 and 84 days. cr = k · w a · t · d , (2) where cr is the corrosion rate in mm year−1, k is a constant in mm year−1 that determines the resulting unit of the cr value (in our case 8.76 · 104), w is a mass loss in g, a is an area of a sample in cm2, t is a time of exposure in h and d is the density of materials in g cm−3. 2.2.2. the method for evaluation of the plastic-pcm compatibility the criterion for evaluating plastic-pcm compatibility is the percentage change in mass of the tested samples. this parameter is determined by calculation according to the gravimetric method defined in the international standard [8] iso 175:2000. ∆m = (mt − mt0 ) mt0 · 100, (3) where ∆m is the percentage change in mass in %, mt is the weight of the sample after removal from pcms in mg and mt0 is the initial weight of the sample before immersion in pcms in mg. 3. results and discussion the material compatibility of the selected four pcms with the plastics and metals was verified by a longterm experiment. the samples of both metals and plastics were immersed in the pcms for the same period with a difference in their extraction frequency. metals were extracted three times throughout the experiment, plastics every 7 days. this time cycle is similar by moreno [3] and óro [2]. in both cases, the evaluation procedure included the visual inspection of the surface changes and the weight changes of the individual samples before and after removal from the pcms, followed by calculations of the evaluation criteria. after removing the metal samples from the pcm, it was found that the samples immersed in inorganic pcm showed visual changes on their surface compared to the samples immersed in organic pcm. all metal samples were tarnished as early as one week after immersion, more pronounced changes were discovered on the surface of the copper samples. the results of the test for metal-pcms pairs are shown in table 4. the table shows only the representatives of the two groups whose weight changes were higher compared to the other material in the group. the evaluation criterion is the corrosion rate that was calculated according to equation (1) and (2). in general, the samples immersed in the organic pcm group showed smaller weight changes and lower cr values (see table 4) than samples immersed in the inorganic group. these results are consistent with our initial assumption and also with the conclusions presented in the previous studies [3, 4]. the most significant mass loss occurred in the samples immersed in pcm rubitherm sp25 (inorganic pcm), the highest values were achieved in the case of the copper, where the cr exceeded 5.4 mg cm−2 year−1 with the value mass loss of 4.4 mg. the cr curves of metal samples immersed in rubitherm sp25 are shown in figure 1b. the cr curves of all metal samples in inorganic pcms (figure 1b) are very similar to those in organic pcms (figure 1a). the difference is in the scale of values, in the case of samples in inorganic pcms (figure 1b), the cr values are much higher. 480 vol. 38/2022 durability of latent heat storage systems (a). linpar 17 (organic pcm). (b). rubitherm sp25 (inorganic pcm). figure 1. comparison of the dependence of cr value (in mg cm−2 year−1) of metals in two types of pcms. figure 2. comparison of the dependence of cr value (in mg year−1) of metals in rubitherm sp25 (inorganic pcm). the cr parameters of the metal samples were compared with the limit values in the guide for corrosion weight loss (in table 3), unlike other study [2], all considered metal-pcm combinations can be classified as recommended for long term service [8]. figure 2 shows the cr values in mm·year−1 from which we determined the estimated service lifetime of these metal samples, which is 100 years for copper, up to 150 years for aluminium, and up to 200 years for brass when immersed in inorganic pcms. this (a). linpar 1820 (organic pcm). (b). rubitherm sp22 (inorganic pcm). figure 3. comparison of the dependence of ∆m value (in %) of plastics in two types of pcms. evaluation will be complemented by a further experiment using the planned interval test according to the methodology in astm [8]. the courses of the values of percentage change in mass of the samples are shown in figure 3, it is a criterion for the evaluation of the mutual compatibility of the plastic-pcm pairs and the calculation is described in section 2.2. the figures are presented for the representatives of the organic and inorganic pcm group as in the case of the metal samples. the results show that samples immersed in the organic group of pcms achieved higher weight changes, where the maximum value of mass change reached 7 %. the organic pcms have a higher ability to penetrate the matrix of the tested plastic samples, the penetration of organic pcms into the plastics materials is demonstrated in figure 3a. in terms of the selected materials, the smallest weight changes were found for pvc-u material and the highest for pp-h immersed in linpar 1820 (organic pcms), similar conclusions are described by castellón [9]. all samples were checked after removal from the pcms and no visual changes were observed on the sample surface, the same results are documented by oró [2]. 4. conclusions in this study, we focused on testing the compatibility of selected plastics and metals with organic and inorganic groups of pcm. according to the results presented in section 3, a significant difference in the weight changes of the samples was found depending 481 s. bantová, m. ostrý, k. struhala acta polytechnica ctu proceedings on the selected group of pcms. it was verified experimentally in the case of immersion of plastics in the inorganic group of pcms, that less mass change and lower penetration of pcms to the matrix of the tested samples occurred. the pvc-u samples did not show any significant weight changes, while a rapid penetration of pcms into the samples of pp-h and pe-hd was clearly visible during the first 4 weeks, but this increase slowed down in the following weeks. the interaction between pcms and metals was verified using the cr value. compared to the plastics, the metal samples in the organic group of pcms showed less weight and visual changes on their surface. the highest values were achieved in the case of the copper immersed in rubitherm sp25, which is the worst variation of the metal-pcm pair in the experiment. the test method for metals used in the experiment was based on the assumption of the normal (constant) corrosion rates during the exposure period of the metals in corrosive environments. in further research, we would focus on the testing metal-pcms combination using the “planned interval test”. this method measures the variation of corrosion rates over time throughout the experiment. acknowledgements this research was funded by czech science foundation, grant number 19-20943s “compatibility of plastics and metals with latent heat storage media for integration in buildings”. references [1] astm g1–03 standard practice for preparing, cleaning, and evaluating corrosion test speciments, astm international, west conshohocken, pa, usa, 2019. [2] e. oró, l. miró, c. barreneche, et al. corrosion of metal and polymer containers for use in pcm cold storage. applied energy 109:449–453, 2013. https://doi.org/10.1016/j.apenergy.2012.10.049 [3] p. moreno, l. miró, a. solé, et al. corrosion of metal and metal alloy containers in contact with phase change materials (pcm) for potential heating and cooling applications. applied energy 125:238–245, 2014. https://doi.org/10.1016/j.apenergy.2014.03.022 [4] g. ferrer, a. solé, c. barreneche, et al. corrosion of metal containers for use in pcm energy storage. renewable energy 76:465–469, 2015. https://doi.org/10.1016/j.renene.2014.11.036 [5] a. j. farrell, b. norton, d. m. kennedy. corrosive effects of salt hydrate phase change materials used with aluminium and copper. journal of materials processing technology 175(1):198–205, 2006. https: //doi.org/10.1016/j.jmatprotec.2005.04.058 [6] p. rolka, j. karwacki, m. jaworski. compatibility tests between three commercially available organic pcms and metals typically used in fin-and-tube heat exchangers. materials 14(18):5172, 2021. https://doi.org/10.3390/ma14185172 [7] astm g 31 standard practise for laboratory immersion corrosion testing of metals, astm international, west conshohocken, pa, usa, 1999. [8] iso 175:2010 plastics – methods of test for the determination of the effect of immersion in liquid chemicals. [2019-02-15]. https://www.iso.org/standard/55483.html [9] c. castellón, i. martorell, l. f. cabeza, et al. compatibility of plastic with phase change materials (pcm). international journal of energy research 35(9):765–771, 2011. https://doi.org/10.1002/er.1723 482 https://doi.org/10.1016/j.apenergy.2012.10.049 https://doi.org/10.1016/j.apenergy.2014.03.022 https://doi.org/10.1016/j.renene.2014.11.036 https://doi.org/10.1016/j.jmatprotec.2005.04.058 https://doi.org/10.1016/j.jmatprotec.2005.04.058 https://doi.org/10.3390/ma14185172 https://www.iso.org/standard/55483.html https://doi.org/10.1002/er.1723 acta polytechnica ctu proceedings 38:478–482, 2022 1 introduction 2 materials and methods 2.1 materials 2.2 methods 2.2.1 the method for evaluation of the metal-pcm compatibility 2.2.2 the method for evaluation of the plastic-pcm compatibility 3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0098 acta polytechnica ctu proceedings 34:98–104, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague phase-field damage of laminated glass in explicit dynamics jaroslav schmidt∗, tomáš janda czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 praha 6, czech republic ∗ corresponding author: jaroslav.schmidt@fsv.cvut.cz abstract. laminated glass is aesthetic but also functional modern material, which has spread from the automotive industry to the civil engineering in the form of design as well as structural elements. proper design of this material requires understanding not only its intact phase, but also its post-breakage behavior. to describe crack initiation and propagation the phase field damage model seems to be suitable tool despite the fact that this approach requires a fine mesh to sufficiently interpolate sharp discontinuity. in this contribution we investigate possibilities of explicit phase field dynamic model with specific use on glass plates, especially applied for dynamic impact load. the dependence of the result on residual stiffness term is examined and incorrect crack speed is observed. usability of phase-field model is still uncertain. keywords: phase field damage model, laminated glass, dynamic damage, glass fracture. 1. introduction the so called phase-field damage model firstly presented in [1] is nowadays very popular approach and it is widely studied. the disadvantage of such a model is the need for a fine mesh to simulate brittle failure, but the main advantage remains the prediction of crack initiation and branching without additional ad hoc criteria. these phenomena are predicted within a variational principle. despite the fact that it was academicly investigated in dynamic regime [2], there is still a few articles focused on application on real structures. this paper focuses on such an analysis by modeling the cracking of glass under dynamic loading. the dynamic model including impact is briefly presented in section 2. the phase-field model formulation follows partially from [3] and [4]. for comprehensive overview see [5]. the implementation details are summarized in section 3 followed by fem model description in section 4. the first numerical results are benchmarked in section 5 and finally the main results are presented in section 6 and concluded in section 7. 2. phase-field model we suppose the elastic damageable material occupying space ω is loaded by external traction forces t on part ωt and the current body deformation is described by displacement field u(x) and damage field d(x). the latter one represents the relative measure of damage as a transition of the intact material d = 0 to a fully developed crack d = 1. for dynamical systems, the common principle of minimum potential is replaced by hamilton’s principle of stationary action and fields u(t) and d(t), t ∈ 〈0,t〉 are obtained by minimization of action a(u,u̇,d) = ∫ t 0 l(u,u̇,d) dt, (1) where l is so called lagrangian which form is postulated as l(u,u̇,d) = k(u̇) − ψe(u,d) − ψs,l(d) + p(u), (2) where k is kinetic energy defined as quadratic form of displacement k(u̇) = ∫ ω 1 2 ρu̇ · u̇ dx. (3) this formulation assumes that kinetic energy is not affected by damage, therefore the functional is dindependent. the second term in equation (2) is elastic potential. it can be additively decomposed on part ψ+ which is degraded by damage through function g(d) and intact part ψ−, i.e. ψe(u,d) = ∫ ω g(d)ψ+(u) + ψ−(u) dx, (4) where ψ• represents elastic energy density of ψ•. in what follows we always assume that g(d) = (1 −d)2 + kres, where kres is small (kres << 1) auxiliary residual relative stiffness which stabilizes the calculations. the third part of (2) is dissipated energy ψs,l which is approximated by following regularization ψs,l = gf cα ∫ ω α(d) l + l|∇d|2 dx, (5) where gf is the critical energy release rate and l is length of the process zone. for brittle material, l can be treated as numerical parameter with value as 98 https://doi.org/10.14311/app.2022.34.0098 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 34/2022 phase-field damage of laminated glass in explicit dynamics small as possible, but not smaller than element size. the function α and associated coefficient cα drives the phase-field model type. two common options are commonly used in literature: α = d with cα = 8/3 and α = d2 with cα = 2 however we restrict to the first one. finally, the last ingredient is external work of traction forces t defined as p(u) = ∫ ωt t ·u dx. (6) minimization of the action integral (1) with additional per partes integration leads to common continous governing equations for displacement field div ( g(d) ∂ψ+ ∂u + ∂ψ− ∂u ) = ρü (7) supplemented by damage-evolution inequality gf cα ( α′(d) l − 2l∆d ) = −g′(d)ψ+, ḋ > 0 (8) gf cα ( α′(d) l − 2l∆d ) > −g′(d)ψ+, ḋ = 0 (9) and approriate boundary conditions. strong-form equations (7)-(9) have a generally unknown solution, therefore some numerical method must be used. 2.1. plate models the model and its properties are determined by the choice of dissipation function α(d), degradation function g(d) and by active/passive elastic energy densities ψ±. four common choices for the latter one are presented in this subsection. despite the fact that we restrict our attention to model p-vd, others are presented. this is because we intend to investigate them in the future research. 3d spectral decomposition it is assumed that material behaves linearly elastically in pre-breakage stage, therefore it can be described by two independent lame’s coefficients λ and ν. since the damage variable starts evolving, the only positive part of energy must be degraded, therefore ψ±(u) = 1 2 λ〈±tr(ε)〉2 + µε± : ε±, (10) where ε is conventional strain tensor defined as symmetric part of displacement field gradient ε = ∇su. positive and negative component of strain tensor follows from the principal strains εi and the dyadic product of eigenvectors pi of ε, thus ε± = ∑ i ±〈±εi〉pi ⊗pi. (11) we denote this model as 3d − sd 3d volumetric decomposition the second common choice of decomposition is volumetric-deviatoric split, which assumes that only compression part of volumetric strain component remains intact, thus ψ+(u) = 1 2 k〈tr(ε)〉2 + µεd : εd (12) ψ−(u) = 1 2 k〈−tr(ε)〉2, (13) where k stands for bulk modolus and εd is deviatoric part of strain ε. this model is denoted as 3d − vd. mindlin-reissner plate besides full 3d model, we investigate the spatialy reduced mindlin-reissner plate model. this model assumes linearly distibuted strain and stress field across thickness and kinematics is driven by middle-surface longitudinal displacement field u, out-of-plane scalar deflection w and vectorial tensor of cross-section rotations ϕ. based on the linearity assumption the longitudinal strain ε and shear strain γ can be established as ε(xs,z) = ∇su(xs) + ∇ssϕ(xs)z, (14) γ(xs) = 1 2 (∇w(xs) + sϕ(xs)) , (15) where z is out-of-plane coordinate meanwhile xs = {x,y} is in-plane coordinate system of middle-surface ωs. the s is auxiliary matrix s = [ 0 1 −1 1 ] . (16) mindlin-reissner plate model ignores normal stress across thickness, therefore in each infinitesimal layer plane-stress situation occurs. therefore stress fields are defined as σ(ε) = dps : ε, (17) τ(γ) = 2νi : γ, (18) where dps is plane-stress elastic tensor defined again through lame’s coefficients λ and ν and i is identity tensor. the last ingredient is thickness integration for spatial reduction. it is impossible to perform it in closed form therefore numerical integration is employed and active/pasive elastic energy become ψ+ = 1 2 ∫ ωs g(d) (∑ i ∆ziεi : σ+i + ξhγ ·τ ) dxs, (19) ψ− = 1 2 ∫ ωs ∑ i ∆ziεi : σ−i + (1 − ξ)hγ ·τ dxs. (20) summation represents dividing thickness into layers i, each represented by in-plane stress σi and strain εi with sub-thickness ∆zi. parameter ξ decides what amount of shear is degraded; ξ = 1 means shear is fully degraded meanwhile ξ = 0 is shear damage-free 99 jaroslav schmidt, tomáš janda acta polytechnica ctu proceedings state. still the stress tensor σ±i must be somehow decomposed. based on this decomposition this model is denoted as p − sd if spectral decomposition is used and is denoted p − vd for volumetric-deviatoric split. the degradation function still remains g(d) = (1 − d)2 +kres, but we found that it is appropriate to select different parameter kres for shear, which is denotes by ksres. 2.2. discrete dynamic model first of all we suppose that the solution is evaluated only in discrete time instants 0 = t0 < t1 < t2 < ... < tn = t selected from the original time interval 〈0,t〉. the action integral (1) is approximated based on this time discretization as n−1∑ i=0 ∆tild(ui,ui+1,di,di+1), (21) where common notation ui = u(ti) and di = d(ti) is used and also ∆ti = ti+1 − ti = ∆t is supposed to be constant. function ld is discretized lagrandian, which approximate the action integral on given subinterval. for purpose of this article, relatively simple expression leading to central difference scheme is used, i.e. ld(ui,ui+1,di,di+1) = l(ui, ui+1 −ui ∆t ,di). (22) as a result the original minimization in continuous time domain fell apart to individual minimizations with respect to each displacement ui and damage di, where i = 0, . . . ,n. the stationarity condition for displacement field ui leads to discrete euler-lagrange equation in form ∂ld(ui−1,ui,di−1,di) ∂ui + ∂ld(ui,ui+1,di,di+1) ∂ui = 0, (23) which can be expressed implicitly in weak form δk(δu) − δψe(δu) + δp(δu) = 0,∀δu (24) where δk = 1 ∆t2 ∫ ω ρ(ui+1 − 2ui + ui−1) · δu dx, (25) δψe = ∫ ω ( g(di) ∂ψ+(ui) ∂ui + ∂ψ−(ui) ∂ui ) · δu dx, (26) δp = ∫ ω t · δui dx. (27) the presented weak form can be further modified to obtained governing equations in strong form, but we employ finite element method for solving spatial distribution of ui which requires weak form as an input. the governing weak form for damage is identical with quasi-static version. the stationary condition δa(u,u̇,d,δd) = 0, ∀δd (28) becomes∫ ω gf cα ( α′(d) l + l∇d∇δd ) + g′(d)ψ+ dx = 0, ∀δd (29) with damage irreversibility condition di(x) ≤ di+1(x), ∀x, i = 0, . . . ,n − 1 (30) this approach therefore induces the following explicit non-incremental dynamic procedure: (i) solve explicitly displacement in time ti by linear problem (24), (ii) solve damage field di by variational ineqaulity problem (29)+(30), (iii) increment time. 2.3. nonlinear hertz force the impact of steel impactor with weight mimp is included in the model through hertz-law. the impactor is characterized by scalar displacement uimp, which is binded to plate by nonlinear contact force f(u,uimp) = k〈uz(ximp) −uimp〉3/2, (31) where ximp is impact position on body ω, uz is out-ofplane component of u = (ux,uy,uz) and k is contact stiffness. although the force is nonlinear the pseudo potential of this force can be still founded. this potential phertz = 2 5 k〈uz(ximp) −uimp〉5/2 (32) is added to external work (6) and it induces force (31) applied to impact point ximp and negative force (31) applied to impactor. additionaly the kinetic energy must be enhanced by kimp = 1 2 mimpu̇impu̇imp (33) with impactor displacement rate u̇imp. this part of kinetic energy is again approximated by discretized lagrangian and discrete euler-lagrange equation induces following additional member of weak form (26), i.e. δkimp = mimp uimp,i+1 − 2uimp,i + uimp,i−1 ∆t2 δuimp (34) 3. implementation for solving spatial distribution of ui, the common finite element method (fem) is used. the displacement field is approximated by nodal values ri through selected basis functions. since the fem is well established, the details of method is not presented here. to handle simultaneously deflection of elastic body and 100 vol. 34/2022 phase-field damage of laminated glass in explicit dynamics l1 = 1.5 m l 2 = 1. 5 m h = 0.02 m e = 72 gpa ν = 0.22 ρ = 2500 kg/m3 x y figure 1. setup of investigated task. impactor position, the all matrices and vectors are extended by one additionaly degree of freedom and final object is marked by tilde. for example displacement vector r̃i or mass matrix m̃ mean r̃i = {ri,uimp}, m̃ = [ m 0 0 mimp ] . (35) with this in hand, the displacement problem (24) after fem assemble process and localization becomes 1 ∆t2 m̃(r̃i+1 − 2r̃i + r̃i−1) + k̃r̃i + f̃ (r̃i) = 0̃, (36) which is linear in parameter ri+1. the matrices k̃ and m̃ are stiffness and mass matrices. presentation of their precise form goes beyond the scope of this paper. the assembling and localization of damage problem (29) leads to linear form hdi + fd = 0, (37) but let us remark several comments. first of all, the discretized damage field di still undergoes irreversibility condition di � di+1 i = 0, . . . ,n − 1, (38) where � operator stands for element-wise comparison. from that reason the linear problem (37) must be solved by nonlinear solver based on variational inequalities to satisfy condition (38). the second remark is about matrices k̃, h and damage-driving force fd. these are ui resp. di nonlinearly dependent and must be assembled in each time steps. it significantly slows down the calculations. 4. setup the all simulations are performed on the same task – unsupported square glass plate hit by steal impactor. this setup follows from real experimental arrangement, see [6]. the plate has dimensions 1.5 × 1.5 m and thickness 0.02 m, but only one quarter is modeled due to symmetry, see figure 1, where material parameters are also presented. the contact stiffness is defined as k = 4 3 √ r 1−ν2 e + 1−ν2imp eimp , (39) where r = 0.05 m is impactor head radius and eimp = 210 gpa and νimp = 0.3. further weight of impactor is mimp = 48.7 kg and initial velocity-free height of impactor is 0.3 m. finally the phase-field parameter l is choosen as element length and gf = (8lf2t )/(3e), where ft = 55 mpa is tensile strength. 5. benchmarks phase-field damage model requires fine mesh with lower-order elements rather than higher-order elements with lower resolution. it is contrary to plate theory, where lower-order elements exhibit shear locking and must be somehow enhanced to overcome this issue. the plate element enrichment leads to additional degrees of freedom which reduces computability and the ability to do repeated calculations. the finer resolution is more important in this analysis than exact shear implementation, therefore suitability of "classic" low-order plate elements is investigated in this benchmark. the first benchmark is behavior of impacted plate in elastic regime. the task setup presented in section 4 is used and the response is calculated in time interval 〈0; 0.01〉 s with time step ∆t = 4 · 10−7 s. resolutions of mesh 50 × 50 and 100 × 100 elements are investigated with two different element typologies: triangles and squares. the reduced integration is employed for shear terms, which enhanced rectangle elements, but triangles still locking. the reason for not using square elements is because they exhibits undesirable zero-energy modes. the results of each case is presented in figure 2, where evolution of impactor acceleration is plotted. this is one of the interesting investigated quantity, which can be experimentaly validated in future research. also the different behavior of two cases stands out more than in impactor displacement function. it is evident from figure 2 that behavior is slighly different for triangles in case of mesh 50 × 50, but both line are comparably the same for case 100 × 100 and shear locking does not cause significant deviation. this statement holds for given time resolution and fraction h/l of course. nevertheless this setup follows from real experiment and we assume that results are adequate. the deflections behaves similarly and do not suffer for larger deviation for mesh 100 × 100. however, the usage of triangle elements exhibit parasitic behavior in the damaged state as is presented in second benchmark. the task is the same as in previous benchmark, but initial crack is incorporated 101 jaroslav schmidt, tomáš janda acta polytechnica ctu proceedings 0 0.2 0.4 0.6 0.8 1 ·10−2 −600 −400 −200 0 time t [s] im p. ac ce le ra ti on [m /s 2 ] 50x50 triangles rectangles 0 0.2 0.4 0.6 0.8 1 ·10−2 −600 −400 −200 0 time t [s] 100x100 triangles rectangles figure 2. evolution of impactor acceleration for triangles (solid) or squares (dashed) mesh plotted for resolution 50 × 50 (red) and 100 × 100 (blue). 0 0.2 0.4 0.6 0.8 0 2 4 6 ·10−3 x coordinate [m] d efl ec ti n [m ] rectangles triangles triangles-full figure 3. deflection distribution on line {(x, y)|y = 0.75}. on plate in this case. the initial crack is implemented through dirichlet boundary condition d = 0.99 on the line {(x,y)|ζ > |x−l1/4|}, where ζ parameter is chosen such that the damage d = 0.99 is enforced at least over whole width of one element. the simplified but still sufficient analysis is performed in this benchmark using so called hybrid approach [7] for phase-field. furthermore, the damage d and updated matrix k̃ from (36) is calculated in each third step to speed up analysis. this approach allows some minor wave propagation through crack, but wave dispersion through crack is still significant. the result also depends on the parameter ξ from (19)-(20). first of all, let us investigate case ξ = 0, therefore shear is not degraded. the results are plotted by solid lines in figure 3, where plate deflection along line {(x,y)|y=0.75} is plotted in time instant t = 0.003 s. the square-topology mesh m1 m2 m3 figure 4. three investigated meshes. behaves reasonable, but triangle-topology mesh does not stop the wave on crack interface. it is caused probably due to element locking. on the other hand, triangles with ξ = 1 behaves almost like squares with ξ = 0 as a result of shear degradation by damage. it significantly reduces the shear locking artifacts. based on these conclusions, the triangle-based mesh with shear reduced by degradation function is used in the next section. 6. results the calculations are time-consuming and therefore only preliminary results of model p-vd are presented in this contribution. the model is still under investigation and there are still many unresolved issues. the same setup, defined in section 4, is used, i.e. the steel impactor hit the thin glass plate. three different meshes m1, m2 and m3 are tested to detect possible mesh-dependency, see figure 4. the model with fully degraded kinematic variables, i.e. ξ = 1.0,kres = ksres = 0.0, does not behave correctly. the impactor hit the plate causing damage localization under impactor. the plate has zero stiffness and impactor flies through fully damaged elements and no longer influences the rest of plate. as a result, the cracks are not evolving further. this behavior is illustrated in figure 5 by solid line, where again the distribution of plate deflection along {(x,y)|y = 0.75} 102 vol. 34/2022 phase-field damage of laminated glass in explicit dynamics 0 0.2 0.4 0.6 0.8 0 1 2 3 ·10−3 x coordinate [m] d efl ec ti n [m ] ksres = 0.0 ksres = 0.05 figure 5. deflection distribution on line {(x, y)|y = 0.75}. (a) (b) (c) (d) figure 6. damage distribution on meshes m1 (a, c) and m2 (b, d) for resolution 50 × 50 (a, b) and 100 × 100 (c, d). the blue represents intact material whereas red represents d close to 1. and in time t = 1.4 ms is plotted. the example is evaluated on mesh m3. the sort of way-out can be shear coupling in damage state, i.e. kres = 0.0,ksres 6= 0. it prevents large slope of deflection curve, see dashed line in figure 5, where ksres = 0.05. the observed behavior is more pleasant and it leads to crack evolving as presented below. moreover the crack is not thickening after additional load. in contrast with that, the case kres 6= 0 causes thickening. mesh dependency the meshes m1, m2 and m3 in two resolutions 50 × 50 and 100 × 100 was used to investigate mesh dependency. the setup is again the same with constant time step ∆t = 2·10−7 s and final time instant tn = 0.01 s. the results of damage fields are plotted in figure 6, where (a) is mesh m1 50×50, (b) is mesh m2 50×50, (c) is mesh m1 100×100 and finally (d) is mesh m2 100 × 100 all evaluated in end time. the mesh m3 behaves similarly as m1 and is not presented here. there is evident that crack topologies are different on coarse mesh, however this discrepancy is removed figure 7. visualisation of plate deflection. after refinement. the slight mesh dependency can still remain in phase-field models for unstable solutions, but it still remains more or less objective approach in contrast with other techniques. time refinement the damage solution (d) in figure 6 and its visualisation of deflection in figure 7 is therefore appropriate and objective without any further improvement during time refinement as seen in the graph 8. the red and blue line are almost identical. an interesting parasitic behavior appears when the residual stiffness is prescribed also to the bending members (kres = 0.01), see black line in graph 8. the residual term causes unwanted harmonic oscillations between impactor and plate. 0 2 4 6 8 ·10−3 −400 −200 0 time t [s] im p. ac c. [m /s 2 ] ∆t = 2 · 10−7, kres 6= 0 ∆t = 1 · 10−7, kres = 0 ∆t = 2 · 10−7, kres = 0 0 0.2 0.4 0.6 0.8 1 ·10−3 −350 −300 −250 −200 time t [s] im p. ac c. [m /s 2 ] detail figure 8. evolution of impactor acceleration for two different time steps (red and blue) and for case kres 6= 0 (black) with detail below. crack speed the main drawback of presented approach is its inability to correctly predict crack velocity. it is evident from near-the-failure evolution of 103 jaroslav schmidt, tomáš janda acta polytechnica ctu proceedings t = 0.00492 s t = 0.00522 s 0.5135 m 0.7210 m figure 9. crack development in two different time instants for mesh m1 100 × 100. crack in figure 9. the crack length increment divided by time increment gives us the numerically predicted crack speed vpf = 692 m/s. however the crack speed of glass is about vreal ≈ 1500 m/s [[8]], which is still twice as much. 7. conclusion the explicit variationally consistent dynamic phasefield damage model was presented in this contribution with additional implementation details and preliminary results. the plate model with volumetricdeviatoric decomposition was successfully implemented, but it still exhibits some issues – especially wrong crack velocity and no crack branching. this two facts may be related, because crack branching is caused by large crack velocity. it is under further investigation as well as implementation of other models (particularly p-sd). acknowledgements this publication was supported by the czech science foundation, the grant no. 19-15326s and by the grant agency of the czech technical university in prague, grant no. sgs21/038/ohk1/1t/11. references [1] b. bourdin, g. a. francfort, j.-j. marigo. the variational approach to fracture. journal of elasticity 91(1):5–148, 2008. [2] m. hofacker, c. miehe. continuum phase field modeling of dynamic fracture: variational principles and staggered fe implementation. international journal of fracture 178(1):113–129, 2012. [3] y. shen, m. mollaali, y. li, et al. implementation details for the phase field approaches to fracture. journal of shanghai jiaotong university (science) 23(1):166–174, 2018. [4] k. pham, h. amor, j.-j. marigo, c. maurini. gradient damage models and their use to approximate brittle fracture. international journal of damage mechanics 20(4):618–652, 2011. [5] j.-y. wu, v. p. nguyen, c. t. nguyen, et al. phase-field modeling of fracture. advances in applied mechanics 53:1–183, 2020. [6] a. zemanová, p. konrád, p. hála, et al. experimental study on the gradual fracture of layers in multi-layer laminated glass plates under low-velocity impact. arxiv preprint arxiv:210712151 2021. [7] m. ambati, t. gerasimov, l. de lorenzis. a review on phase-field models of brittle fracture and a new fast hybrid formulation. computational mechanics 55(2):383–405, 2015. [8] f. barstow, h. edgerton. glass-fracture velocity. journal of the american ceramic society 22(1-12):302–307, 1939. 104 acta polytechnica ctu proceedings 34:98–104, 2022 1 introduction 2 phase-field model 2.1 plate models 2.2 discrete dynamic model 2.3 nonlinear hertz force 3 implementation 4 setup 5 benchmarks 6 results 7 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0038 acta polytechnica ctu proceedings 34:38–42, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the effect of sampling volume size on the apparent stiffness of jointed rock mass martin lebeda∗, petr kabele czech technical university in prague, faculty of civil engineering, thakurova 7, 166 29 prague 6, czech republic ∗ corresponding author: martin.lebeda@fsv.cvut.cz abstract. overall mechanical properties of a jointed rock mass are strongly affected by discontinuities – fractures – that naturally occur in rocks. stochastically-generated discrete fracture network (dfn) modeling, which uses a probabilistic approach to describe the spatial distribution of fractures, such as position, size, or orientation, offers an explicit way to describe geometry of the fracture system. many in-situ measurements and analyses presented in literature indicate that fractures’ sizes can be adequately represented by the power law probability distribution. the parallel plate model of individual fractures combined with an averaging technique makes it possible to estimate the overall compliance or stiffness of jointed rock mass (oda et al. [1]). in the present study, a series of numerical simulations of jointed rock mass modeled by dfn and oda’s approach were conducted to analyze the effect of different sizes of the sampling volume on the overall elastic moduli. the results of the numerical study show that the variance as well as the average of the apparent stiffness decrease as the size of sampling element grows. keywords: jointed rock mass, discrete fracture network (dfn), volume averaging, representative volume element (rve), apparent elastic modulus. 1. introduction brittle structures, such as joints and faults, are found in almost every rock mass body. they vary greatly in quantity, dimensions, and arrangement. the presence of these fractures strongly affects the deformational response of the rock mass to stress and needs to be taken into account in geotechnical calculations. the finite element method (fem) formulated on the basis on the governing equations of solid mechanics is often used for these calculations. in such a case, it is feasible to model discretely (using, e.g., interface elements) only the less-frequent largest fractures, while the smaller ones, which occur in a large quantity, can be represented by the means of so-called equivalent continuum. in the latter approach, a solid containing discontinuities is treated as a homogeneous continuum, whose stiffness or compliance is determined so as to correspond to the relations between the overall stress and overall strain of the fractured body. to this end, homogenization or averaging [2] over a certain sampling volume of the fractured body can be used. the smallest sampling volume element of the fractured body “for which the usual spatially constant ‘overall modulus’ macroscopic constitutive representation is a sufficiently accurate model to represent mean constitutive response” [3], is called the representative volume element (rve). if it is possible to identify an rve that is small enough to be idealized as a material point on the scale of the structural problem analyzed by the fem, then the jointed rock mass can be modeled as a uniform equivalent continuum. if, however, the size of the rve is larger than the resolution required for the structural problem, it is necessary to employ the concept of the statistical volume element (sve). the overall properties of the sve are, again, determined by homogenization or averaging. however, they need to be evaluated over a sufficient number of sampling volumes of the fractured solid and interpreted statistically. the rock mass in the structural problem solved by the fem is then modeled as a nonuniform equivalent continuum. to this end, random fields respecting the statistics of the sves can be used. stochastically-generated discrete fracture network (dfn) models, which uses a probabilistic approach to describe the spatial distribution of fractures, such as position, size, or orientation, offers an explicit way to describe geometry of the fracture system in rock mass. many in-situ measurements and analyses presented in literature indicate that fractures’ sizes can be adequately represented by the power law probability distribution [4], while various distributions for spherical data can model their orientations [5]. one appealing aspect of the dfn modeling is that the parameters of the probability distributions can be identified based on data from structural-geological survey, e.g. [6]. it follows that stochastic realizations of the dfn may serve as the geometrical description of the fractures in rock mass to be used for the earlierdiscussed sampling and determination of the rve or the sve properties [7–9]. in this article, we present results of a numerical case study, in which a body of a jointed rock modeled by stochastic dfn realizations is sampled by volume elements of different sizes and the apparent overall young moduli are evaluated and 38 https://doi.org/10.14311/app.2022.34.0038 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 34/2022 the effect of sampling volume size on the apparent stiffness of jointed rock mass analyzed. the dfns were generated using the power law probability distribution of fractures’ sizes and uniform probability distribution of fractures’ location and orientation. the parallel plate model of fractures and the averaging scheme based on the crack tensors proposed by oda [1] were employed to calculate the overall moduli. in contrast to the previous studies [7– 9] we carried out the analysis in 3d, adopted some different assumptions about the probabilistic distributions of fractures’ sizes and orientations or used a different averaging approach. 2. dfn model of joints’ geometry geometry of individual fractures and the whole fracture system in the modeled rock mass are described by 3d stochastically generated dfn with the following assumptions. 2.1. assumptions on fractures’ orientation, spatial distribution, and shape fractures in natural rock can be usually grouped into sets (populations) based on their geological origin and dominant orientation. however, for the sake of simplicity, to avoid the issue of overall anisotropy of the rock mass, we assume the distribution of fracture orientations as uniform on the sphere (fisher et al. [5]). fractures’ centers are generated by poisson random process, so the probability distribution of center coordinates in space is also uniform. the total number of generated fractures is controlled by a given volumetric density p30 (number of fractures’ centroids over a unit volume). individual fractures are idealized as squares. 2.2. power law distribution of the fracture sizes fractures’ sizes are assumed to follow the power law distribution, whose probability density function is expressed as: p(x) = α− 1 xmin ( x xmin )−α for x > xmin and α > 1, (1) where x is the radius of the circle circumscribed to the fracture, xmin is the minimum fracture size (location parameter), α is the law’s exponent (shape parameter). the parameters used for generating stochastic dfns and for subsequent calculations are listed in table 1. the volumetric density and power law exponent are adapted from our previous work [10], where they were estimated for a highly jointed metamorphic rock based on data from structural-geological mapping of fracture traces on tunnel walls. thus, the range of fracture sizes covered by the model is appropriate for the scale of an underground structure (e.g. tunnel). the dfn models used in this paper were generated using dfram software (kabele et al. [11]). parameter value volumetric density p30 [ 1m3 ] 2.56 power law exponent α [-] 3.4 minimum fracture size xmin [m] 0.3 volume of the rock mass [m3] 100·100·100 table 1. parameters of dfns. 3. the model of the overall stress-strain behavior of jointed rock mass oda [1] proposed the following formula for overall apparent elastic stress-strain relation of jointed rock mass: εij = 1 e [ (1 + ν)δikδjl −νδijδkl + ( 1 kn − 1 ks ) fijkl + 1 4ks (δikfjl + δjkfil + δilfjk + δjlfik) ] σkl = cijklσkl, (2) where e and ν are young’s modulus and poisson’s ratio of the intact rock, respectively, kn and ks are nondimensional parameters related to the fracture (joint) normal and tangent stiffness, respectively, δij is kronecker’s delta, and fij and fijkl are so-called second and fourth rank crack tensors, respectively: fij = 1 v m∑ k=1 s(k)r(k)n (k) i n (k) j , (3) fijlm = 1 v m∑ k=1 s(k)r(k)n (k) i n (k) j n (k) l n (k) m . (4) here v is the volume of the sample, s(k) is the area of k-th fracture inside the sampling volume, r(k) is a typical size of the fracture equal to a diameter of a circle with the same area and n(k)i are the components of a unit vector normal to the fracture. note that the typical size r is assumed as a property of the fracture independent on the size of the sampling volume and it is calculated using the original area of the fracture generated in stochastic dfn. individual fractures are in oda’s approach represented by the so-called parallel plate model as a set of parallel flat plates connected by two springs that symbolize normal and shear stiffness. the fractures’ stiffnesses influence the evaluation of the compliance tensor through the non-dimensional stiffness parameters kn and ks, as seen in eq. (2). the values of the rock’s and joints’ mechanical parameters used in this paper are listed in tab. 2. the joint’s stiffnesses were determined assuming a small stress change in a rock at the depth of 550 m and the other parameters 39 martin lebeda, petr kabele acta polytechnica ctu proceedings parameter value normal stiffness kn [-] 1.356 shear stiffness ks [-] 1.188 poisson’s ratio of intact rock ν [-] 0.164 elastic modulus of intact rock e [gpa] 44.0 table 2. parameters of oda’s parallel plate model. (e and ν) were determined with the basis on in-situ measurements on the same rock as described in [10]. 4. methodology of the analysis 4.1. general considerations we consider that the overall stiffness of the rock mass evaluated by oda’s method on the basis of the dfn would be eventually used for a finite element analysis of stress and deformation in the vicinity of an underground structure, e.g. a tunnel or deep geological repository of radioactive waste. in this scenario, it is plausible to assume that the equivalent continuum model would be employed to account for joints only up to a certain maximum size amax, while larger fractures would be represented as discrete discontinuities with, e.g., interface elements or extended fem. in the context of the underlying dfn it means that only the abundant smaller joints need to be present in the stochastically generated dfn, while the less frequent larger fractures would be mapped and implemented in a deterministic way. to set the threshold value amax, we used a sufficiently low value of the power law’s complementary cumulative distribution function (ccdf), which is defined as the probability of occurrence of a fracture with size greater than amax. with the parameters listed in tab. 1 and ccdf equal to 1·10−5, the corresponding limit fracture size is amax = 51.4 m·51.4 m = 2642 m2. this value seems acceptable also from the geological survey’s point of view, as brittle geological structures of this size are likely to be mapped deterministically by boreholes and trial tunnels. when the dfn is generated by the dfram program, centers of individual fractures are placed by a random process inside a given prismatic bounding box. thus, there is an inner region along the bounding box boundary, which could be intersected by fractures from outside of the box, but these fractures are not generated as their centers lie outside the box. the “thickness” of this boundary zone corresponds to the radius of the circle circumscribed to the largest fracture, that is 51.4 m2 · √ 2 = 36.3 m. to avoid distortion of results by this effect, the sampling volumes must be taken from a subdomain of the dfn bounding box excluding these boundary regions. 4.2. generation of dfn models and sampling volumes to obtain a statistically relevant set of sampling volumes for the evaluation of the apparent overall stiffness, 10 different stochastic realizations of dfn model with the same probabilistic parameters and the same parameters of rock mass (tab. 1 and tab. 2) but different random seed were generated. the bounding box of the dfn fractures’ centroids was a cube with side length of 100 m. cubic sampling volumes of different sizes were created by clipping the dfn models, as shown in fig. 1. the samples were always concentric with the dfn bounding box. considering the dfn box size and the elimination of the boundary effect, the side length of the largest sampling volume was 27.3 m. the edge of the smallest sampling volume was chosen 0.5 m to be slightly less than the diagonal of the smallest fracture (2·xmin). in total, 10 sampling volumes (with edge sizes 0.5, 1.0, 2.0, 3.0, 4.0, 5.0, 10.0, 15.0, 20.0, 27.3 meters) were extracted from each realization of the stochastic dfn. figure 1. sample volume. 4.3. averaging procedure – calculation of overall strain and apparent stiffness averaging procedure – calculation of overall strain and apparent stiffness for each sampling volume, the overall apparent young moduli were calculated using eq. (1). although uniform probability distributions were used for the location and orientation of the fractures in the underlying dfn model, which would suggest isotropy, the apparent moduli were calculated for three orthogonal directions: ex = 1/c1111,ey = 1/c2222,ez = 1/c3333. this way of calculation allowed us to analyze not only the magnitude of the stiffness but also its directional dependence. 5. results figures 2–4 present the dependence of apparent deformation moduli calculated in three orthogonal di40 vol. 34/2022 the effect of sampling volume size on the apparent stiffness of jointed rock mass rections on the size of sampling volume. figure 2. apparent deformation moduli – x axis. figure 3. apparent deformation moduli – y axis. figure 4. apparent deformation moduli – z axis. the results show the same trends for moduli in all three directions. as the sampling volume grows, the mean of the apparent moduli obtained from different realizations of the dfn decreases and converges to a certain value. the maxima have the same tendency, while the minima grow. the dependence of the mean values can be well approximated by an exponential curve, as shown by the dotted line. the graphs also show that the smaller are the sample volumes, the greater are the deviations of the moduli obtained from the different realization of the dfn. 6. preliminary analysis of anisotropy of apparent stiffness from figures 2–4 it is obvious, that in spite of the uniform statistical distribution of fractures’ orientation in the underlying dfn, the apparent overall moduli of the rock mass in three orthogonal directions are not equal. this implies that the overall stiffness of the rock mass predicted by the model is not isotropic. hereafter, we present the results of a preliminary analysis that will allow us to gain some insight into how the detected anisotropy is related to the sampling element size. various measures of anisotropy have been proposed in the literature, especially in connection with crystal elasticity – see a review e.g. in [12]. however, as in the present study we examine only the equivalent deformation moduli, we define the index of anisotropy ad-hoc as the coefficient of variation of the apparent moduli ex,ey,ez evaluated in three orthogonal directions: α = std dev(ex,ey,ez) mean(ex,ey,ez) . (5) it is obvious that for an isotropic material, where all moduli are equal, is zero, whereas anisotropy, manifested by a larger variation among the moduli, results in a larger value of α. figure 5. sample size dependence of the index of anisotropy α. figure 5 shows the dependence of the index on the sampling volume element size. as in the case of the elastic moduli, we can see that with increasing size of the sampling volume, the mean and maximum values as well as the scatter of the index decrease. the minima do not exhibit a clear trend and fluctuate close to zero for all volume sizes. the mean value of α for the largest volume is approaching, but is not equal to, zero. this is understandable, since the index α can be only nonnegative and zero mean value would require that in all model realizations a perfectly isotropic response was achieved. these observations indicate that the model converges to an isotropic response (in the sense of ex ≈ ey ≈ ez). however, we point out that these are only preliminary conclusions and a more systematic analysis involving the complete apparent stiffness or compliance tensor and appropriate index of anisotropy should be carried out. 7. conclusions the concept of using oda’s [1] averaging procedure in conjunction with 3d stochastic dfn has been examined with the aim to determine the dependence of apparent rock mass stiffness on sampling volume size. the results of the case study allow us to draw some interesting conclusions: • with growing sample sizes, the mean value of the apparent stiffness is not constant but decreases and converges to a certain value. the minima and maxima converge to the same value. 41 martin lebeda, petr kabele acta polytechnica ctu proceedings • in spite of assuming directionally and spatially uniform probability distributions of fractures, the power law variability of fracture size is sufficient to produce anisotropic overall deformation moduli. the measure of anisotropy decreases with increasing sampling volume size. • the great variance of stiffnesses as well as anisotropy calculated on smaller sampling volumes (up to about 15 m large side of the sample cube with the parameters used in the present study) means that these volume elements cannot be regarded as representative (rve), but stochastic (sve). the corresponding overall properties must be then considered not as effective, but apparent [13]. should these overall properties be used to define equivalent continuum for finite element analysis of an underground structure, the variance must be taken into account, e.g. by using the concept of random fields. it is noted that this may be relevant for analysis of many structures in underground engineering, such as tunnels, where, due to the spatial variability of the stress field and the size of the structure, model resolution on the order of 100 m or even 10−1 m may be required. • on the other hand, if the required finite element model resolution is on the order of 101 m or larger, using the mean stiffness obtained by the averaging over rve as uniform effective property of the equivalent continuum may be acceptable. this could be the case, for example, when variations of the geostatic stress field need to be estimated for hydrogeological simulations on a larger scale. acknowledgements this paper was financially supported by czech technical university in prague under no. sgs project sgs2/037/ohk1/1t/11. references [1] m. oda, t. yanabe, y. ishizuka, et al. elastic stress and strain in jointed rock masses by means of crack tensor analysis. rock mechanics and rock engineering 26(2):89– 112, 1993. https://doi.org/10.1007/bf01023618. [2] m. hori, s. nemat-nasser. on two micromechanics theories for determining micro–macro relations in heterogeneous solids. mechanics of materials 31(10):667–682, 1999. https://doi.org/10.1016/s0167-6636(99)00020-4. [3] w. j. drugan, j. r. willis. a micromechanics-based nonlocal constitutive equation and estimates of representative volume element size for elastic composites. journal of the mechanics and physics of solids 44(4):497–524, 1996. https://doi.org/10.1016/0022-5096(96)00007-5. [4] e. bonnet, o. bour, n. e. odling, et al. scaling of fracture systems in geological media. reviews of geophysics 39(3):347–383, 2001. https://doi.org/10.1029/1999rg000074. [5] n. i. fisher, t. lewis, b. j. j. embleton. statistical analysis of spherical data. cambridge university press (cup), 1993. https://doi.org/10.1017/cbo9780511564345. [6] p. kabele, o. švagera, m. somr, et al. mathematical modeling of brittle fractures in rock mass by means of the dfn method, 2017. [7] k.-b. min, l. jing. numerical determination of the equivalent elastic compliance tensor for fractured rock masses using the distinct element method. international journal of rock mechanics and mining sciences 40(6):795–816, 2003. https://doi.org/10.1016/s1365-1609(03)00038-8. [8] m. gutierrez, d.-j. youn. effects of fracture distribution and length scale on the equivalent continuum elastic compliance of fractured rock masses. journal of rock mechanics and geotechnical engineering 7(6):626–637, 2015. https://doi.org/10.1016/j.jrmge.2015.07.006. [9] h. ni, w. xu, w. wang, et al. statistical analysis of scale effect on equivalent elastic modulus of jointed rock masses. arab j geosci 9(3):206, 2016. https://doi.org/10.1007/s12517-015-2190-z. [10] m. lebeda. určení tuhosti porušeného horninového masivu s využitím dfn modelu, 2020. [11] p. kabele, o. švagera, m. somr, et al. mathematical modeling of brittle fractures in rock mass by means of the dfn method, 2017. [12] s. i. ranganathan, m. ostoja-starzewski. universal elastic anisotropy index. physical review letters 101(5):055504, 2008. https://doi.org/10.1103/physrevlett.101.055504. [13] t. kanit, f. n’guyen, s. forest, et al. apparent and effective physical properties of heterogeneous materials: representativity of samples of two materials from food industry. computer methods in applied mechanics and engineering 195(33-36):3960–3982, 2006. https://doi.org/10.1016/j.cma.2005.07.022. 42 https://doi.org/10.1007/bf01023618 https://doi.org/10.1016/s0167-6636(99)00020-4 https://doi.org/10.1016/0022-5096(96)00007-5 https://doi.org/10.1029/1999rg000074 https://doi.org/10.1017/cbo9780511564345 https://doi.org/10.1016/s1365-1609(03)00038-8 https://doi.org/10.1016/j.jrmge.2015.07.006 https://doi.org/10.1007/s12517-015-2190-z https://doi.org/10.1103/physrevlett.101.055504 https://doi.org/10.1016/j.cma.2005.07.022 acta polytechnica ctu proceedings 34:38–42, 2022 1 introduction 2 dfn model of joints’ geometry 2.1 assumptions on fractures’ orientation, spatial distribution, and shape 2.2 power law distribution of the fracture sizes 3 the model of the overall stress-strain behavior of jointed rock mass 4 methodology of the analysis 4.1 general considerations 4.2 generation of dfn models and sampling volumes 4.3 averaging procedure – calculation of overall strain and apparent stiffness 5 results 6 preliminary analysis of anisotropy of apparent stiffness 7 conclusions acknowledgements references https://doi.org/10.14311/app.2022.33.0444 acta polytechnica ctu proceedings 33:444–451, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague engineering properties of geopolymer concrete: a review eliana parcesepea, ∗, carmine limab, giuseppe maddalonia, maria rosaria peccea a university of sannio, department of engineering, piazza roma, 21, benevento 82100, italy b items srl, piazza guerrazzi, 1, benevento 82100, italy ∗ corresponding author: parcesepe@unisannio.it abstract. geopolymer concrete (gpc) could be a solution that uses a cementless binder and recycled materials for producing concrete, while reducing the carbon dioxide emission and the demand for raw materials. in addition to the environmental aspect, previous studies on gpc showed that it can achieve mechanical characteristics higher than those of ordinary portland concrete (opc) such as greater strength a few days after casting, and it can be suitable for structural applications. in this paper, the state-of-the-art review of gpc is presented through an extensive literature analysis to determine the most recent information regarding the engineering properties of geopolymer concrete and the critical issues that prevent its widespread use and to put forward suggestions for future research. in particular, the physical properties in both fresh and hardened states and the mechanical characteristics are investigated; the structural performance of geopolymer concrete elements is also outlined. keywords: fresh and hardened properties, geopolymeric concrete, recycled materials, structural properties. 1. introduction the issue of global warming and, more generally, environmental pollution concerns all areas of science, technology and industry, including construction. in particular, the construction field is characterized on the one hand by an enormous demand for energy and on the other hand by the significant production of raw materials. from this point of view, the use of more eco-friendly building materials, such as geopolymer concrete (gpc), can be a valid alternative to ordinary portland concrete (opc). the environmental and economic concerns associated with conventional cement-based building materials have led the scientific and technical community to explore possibilities in the use of alternative materials. currently, the concrete industry faces challenges due to the growing demand for portland cement, such as the increasingly limited limestone reserves and the increase in carbon taxes due to greenhouse gas emissions resulting from the production of opc [1]. these issues require the development of alternative binders, such as alkali-activated cement, with the aim of reducing the environmental impact of buildings, the use of a greater percentage of pozzolan waste and improving concrete performance [2]. from this perspective, geopolymers are alternative binders that attract considerable attention due to their early compressive strength, low permeability, good chemical resistance and excellent fire-resistant behaviour. moreover, geopolymers are ecological materials whose production reduces co2 emissions with respect to opcs [3]. despite the positive results available in the scientific literature, the use of gpcs in practice is still limited mainly due to production costs, the wide variety of properties that can be obtained with different mixes of gpc, the lack of more extensive studies both on material and structural elements and the absence of specific design rules and codes. in this paper, a comprehensive literature review is summarized to assess the critical issues of geopolymer concrete and identify the basic aspects for a research project aimed at an experimental program. the main features of the material are summarized to briefly describing its composition, production methods and current applications. then, the physical and mechanical characteristics of the material are discussed, and the structural performance of the reinforced elements made of geopolymer concrete is also analysed. 2. an overview on geopolymers 2.1. composition the term "geopolymer" is generically used to describe an amorphous alkali-aluminosilicate, which is also commonly used to represent "inorganic polymers", "cements activated with alkalis", "geo-elements", "related ceramics with alkali ", etc. despite the variety of nomenclatures, these terms describe all the materials synthesized using the same chemical process [5, 6]. geopolymers consist of basic and additional components: the former is aluminium silicate (binder), sand, alkali (mortar activator solution) and aggregates (to produce concrete); additional components 444 https://doi.org/10.14311/app.2022.33.0444 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 engineering properties of geopolymer concrete figure 1. classification of geopolymer concrete. figure 2. geopolymer concrete road pavement [4]. can be extra water, plasticizers and fibres. specifically, geopolymer concrete can be produced by polymerizing aluminosilicates such as fly ash (fa), metakaolin (mk), slag (sg), rice husk ash (rha) and wood ash with high calcium content (hcwa) by activation with alkaline solution. figure 1 shows the composition of the raw materials at the base of geopolymer concretes (approximately 80% of the composition) and the activators used (approximately 20% of the composition). various alumina silicate resources have been adopted for producing geopolymers, but from the analysis, fa is the most commonly used and widely tested resource used for geopolymer concrete, showing a high reactivity and more durability than metakaolin-based geopolymers. however, slagbased geopolymers are considered to have high early strength and greater acid resistance than metakaolin and fly ash-based systems [3, 7]. these aspects show that the microstructures and chemical properties of geopolymers vary widely, although many physical properties of geopolymers prepared from various aluminosilicate sources may appear to be similar. it would be important to identify more standardized mixtures, while considering economic aspects, to be able to define the mechanical characteristics with reliable formulae and provide design procedures, also achieving classification. 2.2. research and applications scientific interest in the field of geopolymers has considerably increased since 2016 [8]; however, geopolymer concrete has not yet obtained international acfigure 3. prefabricated geopolymer concrete beam [4]. ceptance as a building material mainly because the production cost of geopolymer concrete is not competitive. furthermore, reliable data are needed on the practicality of using geopolymer concrete as a structural reinforcement element to develop design procedures. the economic aspect is certainly very influential on the diffusion of the material due to the transport system, which is not as effective as that of cement. mathew et al. [9] estimated that the cost of gpc based on coal ash and granulated blast furnace slag (ggbs) can be more than twice that of opcbased concrete if the difference in the rate of transport is considered; otherwise, normalizing this effect, the cost difference between the two types of concrete is equal to 7%. however, the advantages in terms of ambient temperature processing, low carbon dioxide emissions, environmental friendliness, carbon dioxide reduction targets and reutilization of waste must be considered; therefore, a procedure to assess the social cost of this material needs to be developed. the geopolymer industry is becoming established, and an increasing number of geopolymer supplier companies are becoming active based on research activities. several applications regard the transportation sector in the usa due to the short setting time of geopolymer cement, which makes it an ideal solution for repairing highways and airport runways (figure 2). furthermore, gpc is now used in australia both for prefabricated elements and in situ casting (figure 3) [4]. 445 e. parcesepe, c. lima, g. maddaloni, m. r. pecce acta polytechnica ctu proceedings research density csa stsb fsc ymd poissonratio activator/ binder ratio curing temperature and time [kg/m3] [mpa] [mpa] [mpa] [gpa] [−] [−] [−] [10] 2330-2430 30-80 3.8-6 5-1 23-31 0.12-0.16 0.35-0.4 60-80 ◦c for 24 h [11] 1876-2555 65-77.9 2.8-5.1 nre 11.2-41.2 0.15-0.19 0.4-0.65 60 ◦c for 24 h [12] 2074-2185 24.9-82.5 1.2-4.3 3.4-5.4 8.2-22.7 nre 0.82-0.92 nre [13] 2400 46.3-57.2 3.1-4.5 nre 17.1-30.8 0.16-0.21 0.4-0.6 60-80-120 ◦c for 4-6-8-10 h [14] 2400 16.2-52.6 2.9-8.4 nre nre 0.13-0.18 0.4 22 ◦c till testing and 50 ◦c for 48 h [15] nre 29-43.5 nre 6.86 10.7-18.4 nre 0.4-0.55 85 ◦c for 20 h [16] nre 17.7-37.9 3.2-5.4 2.9-4.1 nre nre nre 60 ◦c for 48 h [17] 2147-2408 47-56.5 2.8-4.1 4.9-6.2 23-39 0.23-0.26 0.45-0.59 23 ◦c till testing a cs: compression strength. b sts: splitting tensile strength. c fs: flexural strength. d ym: young modulus. e nr: not reported. table 1. properties of geopolymer concrete. 3. mechanical properties of geopolymer concrete 3.1. fresh and hardened properties extensive studies have been performed to assess the performance of geopolymer concrete, showing that the behaviour is affected by the following parameters: • chemical composition and mineralogy of the precursor materials. • quantity, shape (solid, liquid) and type of alkaline activator (optimal na/al ratio: 1-1.5). • si/al ratio of the precursor materials (optimal: 2.5-4) and quantity of the available calcium source, such as portland cement, blast furnace slag and lime. • total water/solids ratio (precursors + alkaline salts). • time and temperature of curing. regarding the precursor materials, mixtures with fine particles activated with alkali exhibit greater workability and require less water due to the reduction in porosity and to the increase in the surface of the finer particles [18, 19]. the molarity of the activator, the superplasticizer and the water content strongly influence the workability of geopolymer concrete [20, 21]. in particular, the addition of naphthalene and polycarboxylate superplasticizing additives increases the workability [22, 23], but the presence of superplasticizers in contents greater than 2% is responsible for a slight degradation of the compressive strength [24]. therefore, the chemical process, called "geopolymerization", depends on the environmental conditions that occur during the process. the existence of an optimum curing temperature (60◦c) that allows the achievement of the best physical and mechanical properties was observed. the stress-strain behaviour of concrete provides insight into its ability to ensure an adequate degree of safety and serviceability in structural applications. different mixes of geopolymer concrete (gpc) were analysed by various authors, and a summary of the engineering properties is reported in table 1. it is noted that conditioned curing at a high temperature for at least 24 h is generally necessary to obtain mean values of compression strength greater 446 vol. 33/2022 engineering properties of geopolymer concrete figure 4. comparison between eurocode 2 and experimental formulae: (a) splitting tensile strength, and (b) elastic modulus. than 50 mpa; therefore, only prefabricated elements can be made that require a large space with a conditioned environment. therefore, it is important to focus research on mixtures and curing techniques but also on the modulus of elasticity that has a scattered correlation with compression strength. 3.2. experimental relationships and standards to establish correlations within the mechanical properties of geopolymer concrete, several attempts have been made from experimental results by various authors. sofi et al. [24] found that the splitting tensile strength and flexural strength of gpc were comparable to those models presented by the australian standard (as 3600) for opc. regarding the stiffness, the modulus of elasticity of gpc was found to be lower than the values predicted by the standards, particularly by aci318 for opc concrete [25]. ryu et al. [6] compared the formulae proposed by aci 363r-92 and model code 1990 with the experimental results. additionally, in this case, the splitting tensile strength (fct) with respect to the cylindric compressive strength (fct) is lower than that provided by the formulae proposed by the standards; therefore, the following equation (equation 1) is proposed: fct = 0.17 (fc) 3/4 (1) other formulations were suggested by lee-lee [25] for both the splitting tensile strength fct (equation 2) and elastic modulus ec (equation 3) compared to the compressive cylindric strength fc: fct = 0.45 (fc) 1/2 (2) ec = 5300 (fc) 1/3 (3) comparing the proposed formulations for the gpc with those reported in eurocode 2 for strength classes of concrete from c12/15 to c50/60 (figure 4), it is noted that also in this case, the european standards overestimate the properties. in particular, the considerable difference concerning the elastic modulus can lead to problems in the serviceability state. therefore, it is necessary to study the possibility of increasing the ratio between the elastic modulus and resistance in mixtures by analysing which components have the greatest influence on the elastic modulus in detail. in the case of tensile strength, it is clear that performance worsens considerably with the increase in compressive strength; therefore, attention must be focused on the analysis of the microstructure and the inert-matrix interface for higher strengths. 4. behaviour of the structural elements 4.1. reinforced concrete beams and columns the research on geopolymer concrete has been extended to structural elements such as beams and columns. studies on fa-based geopolymer concrete have shown that the behaviour of gpc beams is similar to that of ordinary reinforced concrete beams, but the predictions from the standards have led to conservative designs [29]. even for gpc columns, insignificant differences were found with respect to ordinary columns. the failure modes observed in the compression tests show the opening of the cracks and the crushing of the concrete in the central section and the buckling of the reinforcement bar in the compression zone [28]. it is well known that crisis due to bending generally occurs with yielded reinforcement and therefore is not substantially affected by the constitutive bond of the concrete. therefore, tests should be carried out above all on beams that fail in shear because the shear mechanisms are substantially influenced by the tensile strength of the concrete and the interlock characteristics of the cracks, i.e., also by the inert/matrix behaviour. furthermore, the more relevant differences in the flexural behaviour of rc beams occur under serviceability conditions that are 447 e. parcesepe, c. lima, g. maddaloni, m. r. pecce acta polytechnica ctu proceedings research variable remarks beam [26–28] reinforcement ratio (ρ) flexural strength and ductility influenced by ρ, similar behaviour to opc beams also in shear, applicability of the calculation methods of reinforced opc beams [29] quantity ofrecycled aggregate greater number and width of cracks, but greater deformation and ductility [30] reinforcement configuration higher shear strength of gpc beamsthan equivalent opc beams [31] steel fibre volume (ω) increase in the cracking loadand the ultimate strength as ω increases column [32] compression strengthof concrete higher ultimate strength, ductility and stiffness of gpc columns than equivalent opc columns [33] steel fibre volume the addition of steel fibres increasesthe bearing capacity up to 56% [30] compression strengthof concrete, reinforcement ratio crushing breakage of a fragile type similar to reinforced opc columns [34] effect of the confinement increase in the strength by approximately 30%,increase in ductility [35] eccentricity of the load,slenderness ratio no significant difference between the columns in gpc and opc table 2. summary of the structural performance of gpc elements. influenced by cracking and the elastic modulus; therefore, experimental tests could also be developed on ties to better understand crack propagation, spacing and width. to increase the performance of geopolymer concrete, the effects of additional steel fibres were investigated in beams and columns [31, 33]. furthermore, to increase the load carrying capacity and ductility of gpc columns, confinement can be applied using double layer gfrp wrapping [34]. table 2 shows the summary of previous literature on fa-based geopolymer concrete elements. 4.2. bond between reinforcing steel bars and geopolymer concrete the mechanism of the bond influences the embedded length of the reinforcing bar and consequently the load-bearing capacity of the structural elements, the crack opening and spacing, and the anchorage length. the bond-slip mechanism of the steel-concrete interface of rebar is especially influenced by the concrete strength in tension, which has been commented in par. 3.2. investigations on the gpc bond behaviour are rather limited; however, research was started by sofi et al. [36], who observed a lower influence of the type of fly ash on the bond strength. interface behaviour was investigated in further studies through pull-out tests summarized in table 3. in general, the tests showed that the bond strength between fly ash-based gpc and steel bars is on average higher than that of opc [37]. in some cases, the link between steel and gpc is so strong that the break involved steel bars, while in the case of traditional concrete, pull-out breakage occurred [15]. further tests are necessary to develop formulations for reliable anchorage lengths in gpc. 5. conclusions the paper presents a comprehensive review on the performance of gpc, demonstrating the benefit of using gpc but also the lack of results in many important aspects that influence the structural behaviour and the reliability of the properties, which does not allow us to assess reliable formulations for design. observations that can address further research are summarized as follows: • the development of the mix proportion of gpc is more difficult than that of opc due to a range of parameters being involved in the matrix of 448 vol. 33/2022 engineering properties of geopolymer concrete research bond strength[mpa] variables [36] 10.5 − 14.7 amount and type of fly ash, bar diameter [38] 14.5 − 35.6 bar diameter, concrete resistance [37] 24.1 − 31.9 curing time [39] 12.7 − 16.6 bar diameter, embedded length, steel fibre volume [40] 3.1 − 6.7 smooth bars: bar diameter, cover/bar diameter ratio, embedded length table 3. summary of bond strength research. geopolymer concrete; therefore, more accuracy in the choice of components is needed. most of the research work is limited to heat curing conditions; however, geopolymer products can spread, especially if they can suitably be developed under ambient curing conditions. therefore, the variation in the properties of ambient temperature cured materials and geopolymer concrete prepared under field conditions needs to be investigated to achieve classifiable products with reliable standard properties. • gpc has considerable potential for use as a construction material, especially due to its very high strength. however, the mechanical properties (elastic modulus and tensile strength) increase less than those of opc with the same compression strength; therefore, the serviceability performance could be the key issue of the design. furthermore, research is lacking on long-term mechanical properties (creep and shrinkage), permeability and overall durability issues that have to be analysed by microstructural analysis. • experimental studies on structural elements show gaps in tests on serviceability conditions, especially in crack propagation. therefore, further research in the fields of multiaxial stress states (efficiency of confinement) and stiffness degradation is needed. • the economic issue is certainly a challenge to face. a solution to reduce the cost may be performed with raw materials regarding the manufacturing process of sodium hydroxide or replacing the fine aggregates with alternative materials such as crusher dust; however, the impact of using such materials on the strength of concrete has to be studied. • the currently available standards are not exhaustive both for the material and for the structural elements; therefore, the unavailability of adequate manufacturing and design provisions represents the main challenge for the mass use and diffusion of geopolymer concrete. in conclusion, the literature analysis revealed that the mechanical properties of geopolymer concrete are affected by curing conditions, and its manufacturing process requires a proper mixed design. therefore, there is an urgent need to develop a user-friendly geopolymer concrete design procedure to carry out a more efficient analysis of the experimental results considering that the material used is comparable for all tests and theoretical models. this is the first step to assess design formulations with adequate safety factors for design procedures. references [1] s. b. singh, g. ishwarya, m. gupta, et al. geopolymer concrete: a review of some recent developments. construction and building materials 85:78-90, 2015. https: //doi.org/10.1016/j.conbuildmat.2015.03.036. 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[40] e. s. dewi, j. j. ekaputri. the influence of plain bar on bond strength of geopolymer concrete. aip conference proceedings 1855:030017, 2017. https://doi.org/10.1063/1.4985487. 451 https://doi.org/10.1155/2014/312139 https://doi.org/10.1016/j.conbuildmat.2017.03.014 https://doi.org/10.1007/s10853-006-0534-5 https://doi.org/10.1016/j.cemconres.2015.02.016 https://doi.org/10.4028/www.scientific.net/amm.578-579.441 https://doi.org/10.1680/macr.14.00125 https://doi.org/10.1063/1.4985487 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0509 acta polytechnica ctu proceedings 38:509–515, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague verification of window properties after 10 years of exploitation: results of measurements in the pavilion laboratory and the climate chamber marek bartko∗, pavol ďurica university of žilina, faculty of civil engineering, department of building engineering and urban planning, univerzitná 8215/1, 010 26 žilina, slovakia ∗ corresponding author: marek.bartko@uniza.sk abstract. the article will deal with the analysis of measured data on a plastic window with thermal insulating triple glazing, which is suitable for low-energy or passive houses. the window was installed in 2011 in the test laboratory of the department of building engineering and urban planning, faculty of civil engineering, university of žilina (slovakia), where it was tested under standard indoor climate conditions and real outdoor climate conditions. surface temperatures on the frame friezes and glass system and heat flux density were recorded at a five-minute time step. in 2020, the window was removed from the laboratory and subsequently tested in a climate chamber. this paper will present the results of these measurements in terms of heat flow density waveforms, heat transfer coefficient, and total solar transmittance through the glazing. subsequently, a simulation model of this window will be created in the environment of a computational program and its verification based on the measurements will be carried out. a series of calculations will be performed on the tuned model and analyses of the results and comparisons will be presented under the same climatic conditions as during the real measurements recorded by the meteorological station. keywords: window, triple glazing, climate chamber, pavilion measurement, building physics. 1. introduction windows are an integral part of the envelope of almost all buildings. they are the most used transparent system in building envelopes. the main function of windows is to protect the indoor environment from the external climate. other important functions are daylighting and insolation of the space [1], fresh air supply by natural ventilation, contact with the outdoors, etc. with increasing demands for thermal protection, window constructions are constantly evolving. physical criteria such as thermal-technical, acoustic, lighting, and hygiene are important. another important aspect of the proper functioning of windows is their installation in envelope construction. window constructions are the most problematic point of the building envelope in terms of thermal insulation and construction design. therefore, continuous research and development of window constructions are essential. a significant problem that needs to be addressed more than in the past is overheating in summer [2]. another problem is the elimination of thermal bridges in window construction. for example, by using triple glazing we limit the formation of water vapor condensation at the bottom of the glazing. in slovakia, thermal protection is dealt with in stn 730540:2019 [3], which follows the ec energy performance of buildings directive (epbd) 2010/31/eu. valid values for window constructions are uw = 1.0 w/(m2 k) from 2016 and from 2021 the recommended value is uw = 0.65 w/(m2 k). since 2011, three different windows have been monitored in the laboratory of the department of building engineering and urban planning. it is a so-called pavilion laboratory. the indoor environment is controlled by a heating system and an air conditioning unit. the outdoor environment is represented by the real climate, which is recorded by a weather station [4]. from the research conducted so far, several results related to temperature, heat flux, and u-value measurements have been published. comparisons have been made with the model in the fem therm software and result from different years were analyzed. during autumn 2020, the laboratory building envelope was renovated and insulated with a new etics system, while one of the measured windows was dismantled and taken for measurement in the climate chamber [5, 6]. a comparison of the measured results from the pavilion laboratory and the climate chamber after ten years of window exploitation at the same boundary conditions is described in this paper. 2. methods of measurement the analyzed plastic window has been installed in the department testing laboratory since 2011 (figure 1). the window was oriented to the south with a slight inclination to the west (15°). from the exterior side, the window was exposed to real outdoor climate conditions, which are monitored and recorded by a mobile experimental weather station [7], located on the roof of a nearby building. the indoor climate is provided 509 https://doi.org/10.14311/app.2022.38.0509 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en marek bartko, pavol ďurica acta polytechnica ctu proceedings figure 1. external and internal view of the built-in plastic window in the test laboratory. material number of glazing gas total solar energy uw uf ug chambers transmittance g [%] [w/(m2 · k)] [w/(m2 · k)] [w/(m2 · k)] plastic 6 triple ar 36 0.80 1.00 0.50 table 1. properties of the measured plastic window given by the manufacturer. by an air-conditioning unit, which maintains it based on the slovak standard boundary conditions: 20 °c and 50 % humidity [8]. the sensors used for the measurements consist of nicr-ni [9] thermocouples and heat flux density plates (hfp), also equipped with a correction thermocouple (standard 120 × 120 mm) and a half-sized (120 × 60 mm) for window frames and sashes [10]. the monitoring points are the frame, sash, and glazing. the datalogger and both types of sensors are from ahlborn. the data logging interval in this case was five minutes. the heat flux density was only recorded at the glazing, as measuring the heat flux density through the window frame requires different input conditions – the glazing must be replaced by a full panel. for this reason, the degradation of the glazing system will be mainly evaluated in this paper. subsequently, the window was dismantled in 2020 during the reconstruction of the facade of the pavilion laboratory. however, it is not at the end of its life. after dismantling, the window was removed and installed in the climate chamber, where it has been and will continue to be tested and measured. the purpose was to compare the parameters of the window measured in the pavilion laboratory with those measured in the climate chamber. for the measurements in the climate chamber, two variants of indoor climate modeling were used, one with a constant indoor temperature and a measurement with a hotbox, the other without the application of a hotbox and with the same temperature waveform as was recorded in the pavilion. for the measurement of the climate chamber was chosen 1st march 2018. the outside temperature ranges from -3.4 °c on a sunny day to -16.7 °c at night. the indoor air temperature was also used. this method was chosen to compare the behavior of the window under the same boundary conditions in the pavilion laboratory and the climate chamber. instead of thermocouples, pt100 and ntc sensors were used to measure temperature. the hfp used is of similar types to those used in the pavilion measurements. when the window was initially installed in the pavilion laboratory, not all the parameters that we now measure in the laboratory were measured. for this reason, the measured values from the pavilion laboratory and the climate chamber are compared with the values set and declared by the manufacturer. the declared values of the window parameters from the manufacturer are summarized in table 1. 2.1. u-value the heat transfer coefficient u of a structural element describes the amount of heat energy that passes through it from one side to the other per second in a square meter of the area at a constant ambient temperature difference of 1 °c. the relation for calculating the heat transfer coefficient u is shown in equation (1): u = 1 r0 = 1 (rsi + r + rse) = 1 1 hi + 1λ + 1 he = q θai − θae , (1) u heat transfer coefficient [w/(m2 k)], r0 structure resistance to heat transfer [(m2 k)/w], rsi internal surface resistance [(m2 k)/w], 510 vol. 38/2022 verification of window properties after 10 years of exploitation . . . figure 2. detail of the window in the wall of the laboratory. marked positions for comparison of surface temperatures. surface transfer stn 73 0540 glazing glazing glazing coefficient [w/(m2 · k)] pavilion clim. chamber clim. chamber with hotbox without hotbox hi 7.62 10.47 10.92 16.51 he 25 14.56 15.92 15.66 table 2. the surface transfer coefficient is defined in stn 73 0540 and calculated using equations (2) and (3). rse external surface resistance [(m2 k)/w], r thermal resistance of the structure [(m2 k)/w], hi surface transfer coefficient at internal surface [w/(m2 k)], he surface transfer coefficient at external surface [w/(m2 k)], λ thermal conductance [w/(m2 k)]. the surface transfer coefficient can be calculated according to the equations (2) and (3): hi = q θsi − θai , (2) he = q θse − θae . (3) 2.2. solar transmittance for transparent and translucent constructions such as window constructions, besides the thermal quantification, the optical properties, especially the solar transmittance, are equally important parameters [11]. there are several ways, to measure solar transmittance. in our case, measuring devices in the form of two pyranometers were used, one of which was mounted on the interior side of the measured window and the other on the exterior side as part of a mobile meteorological station. from the values measured in this way, the solar transmittance is finally determined as the ratio of the observed solar radiation intensities behind (from the interior side) and in front (from the exterior side) of the measured window located in the pavilion laboratory and subsequently in the climatic chamber. 2.3. window simulation in simulation software in the design and planning process of a building, thermal bridges are analyzed for several construction fragments, including the window frame detail and the window fitting into the opening. the analysis can be carried out using various simulations, however, modeling different window construction is quite timeconsuming [12]. in the field of two-dimensional modeling of heat transfer in buildings, several software is available. this article used therm software, developed by lawrence berkeley national laboratory. the simulation results in the form of surface temperatures were compared with the actual measured surface temperatures in the pavilion laboratory. a detailed view of the modeled window is shown in figure 2. these windows were placed in the same composition as in the laboratory to achieve the best possible temperature match. the simulation was carried out in the lining and the sill. 3. results of measurements for the calculation of the surface transfer coefficient were used equations (2) and (3). results are summarised in table 2. the results show in some cases (especially outdoors) a large difference between the 511 marek bartko, pavol ďurica acta polytechnica ctu proceedings heat transfer specified by the glazing glazing glazing coefficient [w/(m2 · k)] manufacturer pavilion clim. chamber clim. chamber with hotbox without hotbox ug 0.5 0.92 0.94 0.96 uw 0.8 1.12 1.13 1.15 table 3. heat transfer coefficient window uw and glazing ug . figure 3. temperature courses for the centre of glazing. very good match of courses at the night (without solar radiation). measured values and the standard values. the difference could be due to the measurement method itself, a non-stationary state, lower accuracy of the thermocouples, or imperfect contact between the sensor and the surface. in the case of the outdoor coefficient, the non-stationary state of the external environment: solar radiation, wind, and rain. the calculated values of the surface transfer coefficient based on the measured heat fluxes are given in table 2 and the heat transfer coefficient glazing in table 3. the measured surface temperature values at each position are shown in figures 3–6. in this case, the temperature waveforms for a selected time interval were compared (from 20:00 28th february 2018 to 23:59 1st march 2018). although there are 6 (10) positions displayed within the window, only 3 specific positions are compared. in other positions the results are similar. figure 3 shows the temperature waveforms at the center of the glazing from the inside. temperatures were measured using hfp. figure 4 and figure 5 show a comparison of the different internal temperatures. in the first case, the steady – constant air temperature is measured with a hotbox in the climate chamber. in the second figure, the non-stationary temperature was used as the boundary condition in the chamber. figure 6 shows the surface temperatures at the window sash location. in table 4 we can see the solar transmittance calculated as the ratio of the observed solar intensities behind (from the interior side) and in front (from the exterior side) of the measured window. measurements were carried out in the pavilion laboratory and the climate chamber. in table 5 we can see a comparison of the surface temperatures measured in the climate chamber and pavilion laboratory with the surface temperatures from the therm simulation software (figure 7). the boundary conditions in the simulation were set approximately according to the real conditions from the measurements in the pavilion. outdoor temperature -10 °c and indoor temperature 20 °c. 4. conclusion this article deals with the comparison of measured parameters of a plastic window after its ten-year exploitation. the measured parameters were compared only with the data given by the manufacturer. the measured window was installed in the pavilion laboratory where it was exposed to the real outdoor climate. it was later dismantled and used for measurements in the climate chamber with the same boundary conditions set as on the selected winter day in 2018. for comparison, a simulation was also created but in the stationary environment of the therm software. the results showed a very good agreement in the shape of the temperature waveform and for the position in the center of the glazing also in the values where the difference is relatively small. in other positions, there is a different temperature approx. 3–4 °c. daily temperatures in the pavilion laboratory are strongly influenced by solar radiation, wind, and rain. discrepancies in the sill results require further analysis, such as the impact of the masking panel compared to the wall. the glazing heat transfer coefficients calculated from the measured heat fluxes in the pavilion laboratory and the climate chamber show similar values but differ significantly from the values given by the manufacturers. the heat transfer coefficient glazing 512 vol. 38/2022 verification of window properties after 10 years of exploitation . . . figure 4. temperature courses for the bottom of glazing. variant 1 with constant indoor temperature. difference of about 3 °c in surface temperatures. figure 5. temperatures courses for the bottom of glazing. variant 2 with non-steady indoor temperature. difference of about 4 °c in surface temperatures. figure 6. temperature courses for the bottom of windows sash. difference of about 3 °c in surface temperatures. specified by the pavilion climatic chamber manufacturer total solar energy transmittance g [-] 0.36 0.34 0.31 table 4. total solar energy transmittance g. 513 marek bartko, pavol ďurica acta polytechnica ctu proceedings position temperature [°c] therm pavilion chamber p la st ic w in do w 1 o ut do or -1 0 in do or 20 17.3 15.1 15.0 2 16.5 14.9 13.3 3 -9.1 -8.3 -8.9 4 -8.5 -8.3 -9.0 5 17.5 9.9 15.4 6 15.9 12.5 16.9 7 -8.7 -10.6 8 -8.9 -10.6 table 5. comparison of surface temperatures measured in the pavilion and climate chamber with temperatures from the therm simulation software. figure 7. temperature waveforms according to the therm software simulation at the lining and sill location. shows a difference compared to the manufacturer’s values of about 90 %. the values of the total solar energy transmittance compared to the manufacturer’s value is quite similar. the value measured in the pavilion laboratory compared to the manufacturer’s value show a deterioration of only about 5.5 %. for the values measured in the climate chamber, a deterioration of 14 % occurs. based on these findings, we can say that the thermal-technical properties of the plastic window after 10 years of exploitation show deterioration. this deterioration of the parameters is influenced by possible leakage of filler gas from the glazing system, non-stationary conditions, especially from the exterior side (solar radiation, wind, rain), and airflow in the vicinity of the construction. the surface temperatures obtained from the simulation were then compared with the surface temperatures measured in the pavilion and the climate chamber. the results showed a slight overestimation of the temperatures in the lining area, however, there is a large difference between the temperatures in the sill area, which may be due to the airflow in the sill area and the stationary environment of the therm software. acknowledgements the research is supported by the grant project vega no. 1/0673/20. references [1] m. glória gomes, a. j. santos, a. moret rodrigues. solar and visible optical properties of glazing systems with venetian blinds: numerical, experimental and blind control study. building and environment 71:47–59, 2014. https://doi.org/10.1016/j.buildenv.2013.09.003 [2] r. ponechal. increasing thermal mass in low carbon dwelling. procedia engineering 111:645–651, 2015. https://doi.org/10.1016/j.proeng.2015.07.062 [3] stn 73 0540:2019: tepelná ochrana budov. tepelnotechnické vlastnosti stavebných konštrukcií a budov. [4] p. juras, d. jurasova. influence analysis of climate data time-step on the accuracy of ham simulation. matec web of conferences 196:02029, 2018. https://doi.org/10.1051/matecconf/201819602029 [5] p. juras, p. durica, m. bartko. long-term window evaluation: comparison of pavilion laboratory and climate chamber measurement. in p. akimov, n. vatin (eds.), xxx russian-polish-slovak seminar theoretical foundation of civil engineering (rsp 2021), pp. 132–140. springer international publishing, cham, 2022. https://doi.org/10.1007/978-3-030-86001-1_16 [6] k. banionis, j. kumžienė, a. burlingis, et al. the changes in thermal transmittance of window insulating glass units depending on outdoor temperatures in cold 514 https://doi.org/10.1016/j.buildenv.2013.09.003 https://doi.org/10.1016/j.proeng.2015.07.062 https://doi.org/10.1051/matecconf/201819602029 https://doi.org/10.1007/978-3-030-86001-1_16 vol. 38/2022 verification of window properties after 10 years of exploitation . . . climate countries. energies 14(6):1694, 2021. https://doi.org/10.3390/en14061694 [7] p. juras, d. jurasova. outdoor climate change analysis in university campus: case study with heat-air-moisture simulation. civil and environmental engineering 16(2):370–378, 2020. https://doi.org/10.2478/cee-2020-0037 [8] p. juras. comparison of triple glazed windows based on long-term measurement. iop conference series: materials science and engineering 415(1):012020, 2018. https://doi.org/10.1088/1757-899x/415/1/012020 [9] nicr-ni thermocouple t190-0. [2022-04-20]. https://www.ahlborn.com/download/pdfs/kap07/ 05draehted.pdf [10] thermal flux plates. [2022-04-20]. http://www. ahlborn.com/download/pdfs/kap13/wflplatten.pdf [11] m. čekon, r. slávik. total solar transmittance quantifying of transparent insulation building materials based on real climate outdoor measurements. energy procedia 132:243–248, 2017. https://doi.org/10.1016/j.egypro.2017.09.694 [12] p. juras. analysis of window model accuracy and its influence on the results compared to the measurement. matec web of conferences 196:02028, 2018. https://doi.org/10.1051/matecconf/201819602028 515 https://doi.org/10.3390/en14061694 https://doi.org/10.2478/cee-2020-0037 https://doi.org/10.1088/1757-899x/415/1/012020 https://www.ahlborn.com/download/pdfs/kap07/05draehted.pdf https://www.ahlborn.com/download/pdfs/kap07/05draehted.pdf http://www.ahlborn.com/download/pdfs/kap13/wflplatten.pdf http://www.ahlborn.com/download/pdfs/kap13/wflplatten.pdf https://doi.org/10.1016/j.egypro.2017.09.694 https://doi.org/10.1051/matecconf/201819602028 acta polytechnica ctu proceedings 38:509–515, 2022 1 introduction 2 methods of measurement 2.1 u-value 2.2 solar transmittance 2.3 window simulation in simulation software 3 results of measurements 4 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.37.0024 acta polytechnica ctu proceedings 37:24–30, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague modelling of fission gas release in uo2 doped fuel using transuranus code jheffry gonzaleza, b, c, ∗, martin ševečeka a czech technical university in prague, faculty of nuclear sciences and physical engineering, jugoslávských partyzánů 1580/3, 160 00 prague 6 – dejvice, czech republic b royal institute of technology, division of nuclear engineering, kungliga tekniska högskolan, 100 44 stockholm, sweden c grenoble inp – phelma, school of engineering in physics, applied physics, electronics & materials science, 3 parvis louis néel, 38016 grenoble cedex 1, france ∗ corresponding author: jheffrygonzalez@gmail.com abstract. the expected benefits from cr-doped fuel are improved retention of fission gases within the pellets due to its large grain size. to demonstrate this, several experiments have been carried out by halden reactor and studsvik. these experiments are now being used to benchmark several fuel performance codes among them transuranus code. all this as part of a coordinate research project (crp) by iaea named testing and simulation for advanced technology and accident tolerant fuels (atf-ts). this work is introducing a novel fission gas diffusivity model for doped fuel in transuranus code. it is observed the benefits of introducing this new model when comparing to the standard model already existing in transuranus. nevertheless, more work needs to be carried out to fully understand all the phenomena involved in adding dopant in uo2 due to change of thermo mechanical properties. keywords: doped fuel, transuranus, atf, fission gas release, diffusivity. 1. introduction the study of fission gas release, fgr, is an important area within the fuel rod performance analysis that needs to be covered because these gaseous fission products in the uo2 fuel pellet are detrimental to the thermo-mechanical properties of the fuel [1]. fission gas release depends on several mechanisms that occur within the fuel pellet, more specifically within the grain (see figure 1). few of these mechanisms are, for instance, thermal contributions: grain boundary diffusion, bubble migration, trapping and resolution; and there are also athermal contributions: recoil, knock-out and sputtering. figure 1. basic mechanisms of fission gas release [2]. because in this work we are dealing with doped fuel, it is vital to understand how the properties and characteristics of this fuel will impact on the fgr mechanisms. although there are other properties that might need to be investigated this work focuses on the selection of a suitable gas diffusivity model for doped fuel. the selected diffusivity model will be implemented in transuranus code. the new implementation done in the source code will be validated against experimental results performed carried out at the halden reactor (ifa 716 experiment) and the studsvik cladding integrity program (scip-ii). this work is also part of a coordinate research project (crp) by iaea1 named testing and simulation for advanced technology and accident tolerant fuels (atf-ts) and the results will be also compared with other members that are simulating this novel fuel in other fuel performance codes. in section 2, the methodological details of this work are explained, a brief explanation of the experiments to compare to and considerations taken when performing the simulations. section 3 shows the results obtained and their comparison against the tests. finally, in section 4, it will be shown all the conclusions and also recommendations for future improvements of the modelling of this fuel in transuranus. 2. methods the methodology used for the development of this work is as follows i) study the models related to fis1https://www.iaea.org/projects/crp/t12032 24 https://doi.org/10.14311/app.2022.37.0024 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en https://www.iaea.org/projects/crp/t12032 vol. 37/2022 modelling of fission gas release sion gas release in transuranus (i.e. fission gas diffusivity, thermal conductivity, etc.); ii) select the gas diffusivity model from transuranus or introduce a new model; iii) verify that the new implemented model works as expected; iv) identify all the fuel properties and reactor conditions required in the input file to run the simulation (type of fuel and cladding, flux, linear power, temperature of the coolant, grain size, etc); v) data collection of the experimental analysis performed by halden and studsvik; vi) validate the implemented model by benchmarking the results obtain from transuranus with the experimental data. fgr is strictly dependent on temperature; therefore, choosing a good thermal conductivity model for this type of fuel is essential. there is no thermal conductivity model available for doped fuel; however, according to arborelius et al. there is only a small change in thermal conductivity of fresh fuel due to doping and this small difference is expected to be negligible after the burnup increase [3].although it would make sense to use standard undoped uo2 thermal conductivity during simulations the reality is that in this work the correlation used is the one proposed by ohira et al. for for high burnup uo2 pellets [4]. the reasoning for this is explained in section 2.4. transuranus uses a mechanistic model for fission gas behaviour and release from doctoral thesis of [5]. for the intragranular diffusion coefficient, transuranus has different models; nevertheless, the model that is most widely accepted for fission gas diffusivity is from [6] which will be named as a standard during this article. 2.1. selection and modelling of fg diffusivity turnbull model for the single gas atom diffusion coefficient can be seen from equations 1 to 4. def f = d1 + d2 + d3 (1) d1 = 7.6 × 10−10e− 3500 t (2) d2 = 3.22 × 10−16 √ re− 13800 t (3) d3 = 3.22 × 10−16r, (4) where deff is the effective diffusion coefficient, d1 is the intrinsic high-temperature component, d2 is the irradiation enhanced thermal component, d3 is the athermal term and r is the rating in w/gu. because che et al. showed a good agreement with halden fgr experiments when using an empirically tuned enhanced gas diffusion, cooper et al. decided to investigate in deep the influence of dopant on fission gas diffusivity. in that work, cooper et al. proposed a diffusion coefficient for uranium doped fuel [7]. equation 5 shows enhancement of the diffusivity, using the undoped uo2 diffusivity model in equation 1 as a baseline. ddoped = exp ( h1 kb [ 1 t − 1 t1 ]) d undoped 1 + exp ( h2 kb [ 1 t − 1 t2 ]) d undoped 2 + d undoped 3 , (5) where kb is the boltzmann constant. h1, h2, t1 and t2 are parameters derived for the multiscale analytical enhanced diffusivity model (see table 1). dundoped1 , d undoped 2 , and d undoped 3 are calculated with equations 2 to 4. because there are a number of uncertainties when obtaining these parameters cooper et al. examined two cases namely case a and case b. the first case has minimal alterations migration energy and sink strength given by the oxygen potential. in contrast, case b overestimates this migration energy and sink strength based on the experience from [8]. parameter case a case b t1 = t2 [k] 1773 1773 ∆h1 [ev] 0.3198 0.3282 ∆h2 [ev] -0.3345 -0.6998 table 1. parameters for the enhanced diffusivity model derived for cr-doped fuel [7]. now, it is important to understand the distinction (if any) between the models before implementing the proposed fg diffusivity for the doped fuel in transuranus. hence, by plotting the fg diffusion, it is clearly depicted the difference between them (see figure 2). for the model of the fg diffusivity of doped fuel, the two cases proposed by [7] are considered and they are shown case a and case b. figure 2. gas diffusivity in undoped uo2 fuel using turnbull et al. model, and gas diffusivity in cr-doped uo2 fuel applying cooper et al. model. from figure 2, it is observed that from 700 to 1600 k there is a major difference between case a of cooper model and the standard diffusion model, whereas for the case b the range is wider from 500 to 25 jheffry gonzalez, martin ševeček acta polytechnica ctu proceedings 1600 k. the ratio between the doped fuel model and the standard model can be as high as 3.4 and 17.5 for case a and b respectively, at around 1000 k. this ratio difference is significant; however, if we consider the range of temperature from 500 to 2000 k the mean ratio will be for case a a factor of 1.63 and for case b a factor 4.66. 2.2. implementation and verification in this work both cases of cooper model are considered. therefore, it was necessary to have two new versions of transuranus that includes case a and b in which the only difference is the parameters in table 1. the implementation done in transuranus was revised by transuranus developers to ensure that the coding is up to their standards and minor modifications of the code were proposed by them in terms of clarity of the code. as evidenced in figure 2, the value of diffusion coefficient of doped fuel is significantly larger over the range of from 500 to 1600 k, this would mean that if the size of the grain remains the same one would expect a greater fgr when comparing to standard uo2 fuel. hence, for the verification of the modelling, several tests were performed. in these tests the grain size was kept the same but using the new model implemented in tu. table 2 shows the percentage of fission gas release inside the fuel rod and the internal rod pressure difference when using turnbull model and cooper model in an another halden case (ifa 429). test cases parameter model value ifa420cd fgr turnbull 12.34 % cooper 23.17 % pressure turnbull 6.45 mpa cooper 7.27 mpa ifa420ch fgr turnbull 17.79 % cooper 42.95 % pressure turnbull 6.95 mpa cooper 11.18 mpa table 2. model verification using ifa 429 case. for both cases, there is a considerable increase in fgr and pressure when using the diffusivity model for doped fuel. this could confirm that the implementation was done correctly and that the benchmark of the code with experimental data can be carried out. 2.3. experimental data the implemented model in transuranus is validated with experimental data performed by halden and studsvik. from the experimental data and fuel characteristics all the input data required is gathered to run the simulation. some of the input data are unchanged overtime while other data are constantly changing as it is the case of the the power, fast flux and temperature of coolant. for this variable data, an additional processing is needed because transuranus have a maximum number of record for the macro-step that can be included in the input file. this limit is up to 5000, however, the data steps provided is around 160000 for ifa 716 and 24000 for scip-ii. hence, a software called fuel rod analysis toolbox (fratoolbox) is used to merge the quantities of the large number of data sets [9]. 2.3.1. halden experiment ifa 716 the main objective of the ifa 716 experiment is to investigate the effect of dopant concentration and grain size on the fgr [10]. the experiment was carried out between january 2011 and may 2015 for a total operation time of 842 days divided in 15 operation cycles. ifa 716 assembly consisted in six rods each with different type of fuel. few of these characteristics for the ifa 716 case (rod 1 and 2) are reported in table 3; clearly, the main differences between these two fuels are the cr2o3 content and the fuel average grain diameter. moreover, there are some pellets that have a hollow section this only applies for the pellets on the top where instrumentation was inserted to obtain, for instance, the on-line measurement of the fuel temperature center line. figure 3 shows the average linear heat rate measured for the rods 1 and 2 versus time [10]. evidently, rod 1 has a total operation time of 840 days while rod 2 only has over 700 days. this difference is due to a malfunction in rod 2 that caused this rod to be removed from the reactor, in contrast, rod 1 (and the others) remained until the end of the test. the average linear heat rate (lhr) of both rods are roughly the same throughout the experiment; therefore, the average lhr of rod 1 will be used during the simulation in transuranus. figure 3. ifa 716 average linear heat rate measured during the test. 2.3.2. studsvik experiment scip-ii the main objective of scip-ii experiment is to investigate the effect of power ramps on doped fuel with 26 vol. 37/2022 modelling of fission gas release ifa 716 rod 1 ifa 716 rod 2 cladding material zircaloy-4 zircaloy-4 fuel material uo2 + cr2o3 uo2 fill gas he he pellet outer diameter [mm] 9.12 9.13 pellet inner diameter, hollow section [mm] 1.8 1.8 cladding outer diameter [mm] 10.75 10.75 cladding thickness [mm] 0.725 0.725 active fuel length [mm] 399.5 399.4 free volume [m3] 5.8 5.95 fuel density [g/cm3] 10.50 10.55 fill gas pressure [mpa] 1 1 fuel cr2o3 content [ppm] 1580 n/a fuel u-235 enrichment [wt %] 4.9 4.89 fuel average grain diameter [µm] 71 11 table 3. fuel characteristics of ifa 716 rod 1 and 2 simulated in this work [11]. a based irradiation. this based irradiation was carried out between september 2006 and august 2009 at the okg2 reactor divided in 3 cycles. similarly, as the ifa 716 case, the fuel characteristics for the scip-ii case are reported in table 4. rods l707 and l708 are fundamentally the same, there is no major distinction between the two according to the given data. what makes these rods different is the base irradiation that they underwent before being subjected to power ramps in a test reactor. both rods have a total operation time of 1095 days as based irradiation with an average lhr of 14 and 13 kw/m for rod l707 and l708, respectively. this led to a mean bu of 28 mwd/kgu in the case of l707 and 26 mwd/kgu for l708. the ramp test for each rod was carried out differently for both, with rod l707 having a total time of 21.33 h while rod l708 with a total time of 14.32 h. the ramp was hold for 12 h and around 1 minute for rod l707 and l708, respectively. figure 4 depicts the average lhr during the ramp test for each rod. in contrast to the halden experiment, there is no in-situ measurement of fuel center line temperature or plenum pressure which will make the implemented model harder to validate against this experiment. the only measurement that is given comes from the postirradiation examination where the data for fgr is obtained. 2.4. additional considerations the only criteria for selection thermal conductivity model is how close the fuel center line temperature (fclt) obtained from transuranus is to the experimental data. transuranus has a set of thermal conductivity models for different fuels including a correlation for the standard uo2. when using the lwr standard correlation (model already implemented in transuranus) proposed by [12], it results in the figure 8 (top) where evidently there is not a good agreement between this model and the experimental figure 4. scip ii ramp test. data. therefore, exploring other thermal conductivity models is necessary to continue to the validation of the fgr. the model that best fit to the experimental results is surprisingly a correlation used for high burnup uo2. consequently, the correlation for high burnup uo2 pellets found in transuranus is the thermal conductivity model of choice for the study of fgr in this work. in addition to using a thermal conductivity for high burnup uo2, it is considered that the gas diffusivity model for doped fuel is the same independently of the chromia concentration in the fuel because this diffusivity model does not have a correlation as function of the amount of chromia content. 3. results the parameters to benchmark are fuel center line temperature, upper plenum pressure, and fission gas release. however, no all data is available for all the rods tested. for instance, for scip-ii, the only data to compare is the fgr obtained in the post-irradiation examination. 27 jheffry gonzalez, martin ševeček acta polytechnica ctu proceedings scip-ii rod l707 scip-ii rod l708 cladding material zircaloy-4 zircaloy-4 fuel material uo2 + cr2o3 uo2 + cr2o3 fill gas he he pellet outer diameter [mm] 8.67 8.67 cladding outer diameter [mm] 10.05 10.05 cladding thickness [mm] 0.605 0.605 active fuel length [mm] 480 480 fuel density [g/cm3] 10.50 10.55 fill gas pressure [mpa] 1 1 fuel u-235 enrichment [wt %] 3.0 3.0 fuel average grain diameter [µm] 49 49 table 4. fuel characteristics of scip-ii rod l707 and l708 simulated in this thesis. 3.1. benchmark of halden ifa 716 because pressure transducer of rod 2 was damaged, there is no benchmark of this for rod 2. on the other hand for rod 1, the pressure of case a and b for rod 1 is the same during the first 300–350 days of operation and for both cases transuranus predicted well the pressure when compared to the experimental data. however, after 350 days of operation transuranus underestimates the pressure when using diffusivity model case a while when using case b the pressure is overestimated (see figure 5). figure 5. predicted vs. measured rod pressure for rod 1, including the two cases for the diffusivity of doped fuel. figure 7 shows the result of the comparison of the predicted fclt versus the measured fclt taken from the test for rods 1 and 2. it can be observed that for rod 1 (see figures 7a and 7b), there is a good agreement between the experimental and simulation when the burnup is low enough (< 17 mwd/kguo2) while at higher burnup there points that fall out of the established uncertainty zone of ±9 % and yet close to this zone. around 100 % of agreement is found if we accept this uncertainty. in case of rod 2, in figure 7c at low burnup the predicted temperature fall within the uncertainty opposite to what occurs at higher burnup with several data points falling far from this zone. this over-prediction of the temperature, after approximately 400 days of operation, is also observed in cooper et al. work; however, the temperature difference between predicted and measured is larger [7]. finally, for the fission gas release, the given data did not contain the evolution of fission gas release versus operation time; instead, what is available is the fission gas release at different stages of the experiment. nevertheless, this is not true for rod 2, where no data was collected at any point and the rod presented signs of leakage at around 700 days of operation. hence, no puncture was performed to obtain the fission gas release. figure 6 shows the predicted fission gas release for rod 1 for both diffusivity cases, the predicted fission gas release for rod 2 and the measures taken by halden at four different times with their uncertainty for rod 1. despite having a good agreement between measured and predicted for fclt and pressure, the fission gas release shows a wide difference between the measured and simulation. this is more evident for rod 1 case a where the prediction of fgr is 1.6 % while case b gives a better results with 3.8 % at 625 days. the measured value for the fission gas release at the same time was 5.71 % ± 1.36 %. surprisingly, the fission gas release of rod 2 is lower than the rod 1 when using diffusivity case b. figure 6. predicted fission gas release for rod 1 and 2. 28 vol. 37/2022 modelling of fission gas release (a) . rod 1 uo2 doped fuel diffusivity case a (b) . rod 1 uo2 doped fuel diffusivity case b (c) . rod 2 standard uo2 fuel small grain figure 7. predicted vs. measured fuel center line temperature for ifa-716 rods 1 and 2. for burnup lower than 17 mwd/kguo2 (shown as ’5’) and burnup higher than 17 mwd/kguo2 (shown as ’?’). 3.2. benchmark of studsvik scip-ii unfortunately, there is no data to compare fclt or pressure as it was done with ifa-716. table 5 shows the fission gas release obtained during the post irradiation examination and the fission gas release predicted by transuranus using fission gas diffusivity case b. for this experiment, the prediction overestimate the experimental data. this is more evident for the 1 minute ramp (rod l708) where there is almost no release according to pie but the prediction is 2.5 %. rod l707 rod l708 fgr / pie 3.0 % 0.04 % fgr / transuranus 5.8 % 2.5 % table 5. fission gas release comparison at the end of ramp 4. conclusions the fission gas diffusivity model was compared to two different experiments, one for ramp test and the other for long period operation. the results show that there are still uncertainties in the prediction using this model in transuranus. the halden experiment, however, was surrounded by some failures in sensors during the experiment. as a result, it is unclear whether these experimental results can be trusted. similarly, there was a lack of data to make a better comparison against the scip-ii experiment. it will be important to continue the validation, perhaps against other experiments, to draw a final conclusion on this model with the transuranus code. additionally, defining the model exclusively for doped fuel can be beneficial to improving the accuracy of the predictions. acknowledgements this project is financed from the state budget by the technology agency of the czech republic and the ministry of industry and trade within the theta programme. a support from the iaea crp atf-ts t12032 is acknowledged. references [1] m. tonks, d. andersson, r. devanathan, et al. unit mechanisms of fission gas release: current understanding and future needs. journal of nuclear materials 504:300–317, 2018. https://doi.org/10.1016/j.jnucmat.2018.03.016. [2] european commission. transuranus handbook, 2019. [3] j. arborelius, k. backman, l. hallstadius, et al. advanced doped uo2 pellets in lwr applications. journal of nuclear science and technology 43(9):967–976, 2006. https://doi.org/10.1080/18811248.2006.9711184. 29 https://doi.org/10.1016/j.jnucmat.2018.03.016 https://doi.org/10.1080/18811248.2006.9711184 jheffry gonzalez, martin ševeček acta polytechnica ctu proceedings figure 8. fuel center line temperature measured by thermocouples during the ifa 716 test compared to simulated fclt using standard thermal conductivity correlation (top) by hading et al. [12] and high burnup thermal conductivity correlation by ohira et al. [4]. [4] k. ohira, n. itagaki. thermal conductivity measurements of high burnup uo2 pellet and a benchmark calculation of fuel center temperature. in proceedings 1997 international topical meeting on light water reactor fuel performance, pp. 541–549. united states (usa), 1997. isbn 0-89448-616-0. [5] g. pastore. modelling of fission gas swelling and release in oxide nuclear fuel and application to the transuranus code. department of energy, politecnico di milano, 2012. [6] j. turnbull, c. friskney, j. findlay, et al. the diffusion coefficients of gaseous and volatile species during the irradiation of uranium dioxide. journal of nuclear materials 107(2):168–184, 1982. https://doi.org/10.1016/0022-3115(82)90419-6. [7] m. w. cooper, g. pastore, y. che, et al. fission gas diffusion and release for cr2o3-doped uo2: from the atomic to the engineering scale. journal of nuclear materials 545:152590, 2021. https://doi.org/10.1016/j.jnucmat.2020.152590. [8] c. matthews, r. perriot, m. w. cooper, et al. cluster dynamics simulation of uranium self-diffusion during irradiation in uo2. journal of nuclear materials 527:151787, 2019. https://doi.org/10.1016/j.jnucmat.2019.151787. [9] k. lassmann, j. van de laar, a. schubert, p. v. uffelen. ‘mergequantities’: a program for adding axial profiles to a {x,y} data set and for merging quantities. joint research centre of the european commission, 2014. [10] k. fuglesang. in-pile results from the fission gas release mechanisms study in ifa-716 after final unloading. tech. rep., organisation for economic co-operation and development halden reactor project, norway, 2016. [11] m. sobieska. production of six fuel rods for fission gas release mechanisms test rig ifa 716. tech. rep., organisation for economic co-operation and development halden reactor project, norway, 2009. [12] j. harding, d. martin. a recommendation for the thermal conductivity of uo2. journal of nuclear materials 166(3):223–226, 1989. https://doi.org/10.1016/0022-3115(89)90218-3. 30 https://doi.org/10.1016/0022-3115(82)90419-6 https://doi.org/10.1016/j.jnucmat.2020.152590 https://doi.org/10.1016/j.jnucmat.2019.151787 https://doi.org/10.1016/0022-3115(89)90218-3 acta polytechnica ctu proceedings 37:24–30, 2022 1 introduction 2 methods 2.1 selection and modelling of fg diffusivity 2.2 implementation and verification 2.3 experimental data 2.3.1 halden experiment ifa 716 2.3.2 studsvik experiment scip-ii 2.4 additional considerations 3 results 3.1 benchmark of halden ifa 716 3.2 benchmark of studsvik scip-ii 4 conclusions acknowledgements references my title https://doi.org/10.14311/app.2022.36.0231 acta polytechnica ctu proceedings 36:231–236, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague probabilistic safety assessment of koeberg spent fuel pool sibongiseni thabethe university of pretoria, national nuclear regulator, centre for nuclear safety and security, engineering building 2, lynwood road, po box 395, pretoria 0002, south africa correspondence: sthabethe@nnr.co.za abstract. the effective management of spent fuel pool (sfp) safety has been raised as one of the emerging issues to further enhance nuclear installation safety after the fukushima accident on march 11, 2011. sfp safety-related issues have been mainly focused on (a) controlling the configuration of the fuel assemblies in the pool with no loss of pool coolants, and (b) ensuring adequate pool storage space to prevent fuel criticality owing to chain reactions of the fission products and the ability for neutron absorption to keep the fuel cool. in support of regulatory functions, the centre for nuclear safety and security (cnss) seeks to perform confirmatory analysis for all potential accident scenarios that may occur in the koeberg nuclear power plant sfp. probabilistic safety assessment (psa) was done using the systems analysis program for hands-on integrated reliability evaluations (saphire) computer code. we present preliminary psa results of initiating events that lead to boiling and cause fuel uncovering, resulting in possible fuel damage in the koeberg nuclear power plant sfp. keywords: computer code, fuel assemblies, probabilistic safety assessment, spent fuel pool. 1. introduction highly radioactive spent fuel assemblies that are unloaded from the nuclear reactor core are typically stored for a certain period in cooling water pools called spent fuel pools (sfp). the safe storage of these spent fuel assemblies in the sfp is very vital since many radioactive fission products could potentially be released into the environment if a severe accident occurred in the sfp [1]. the fukushima daiichi nuclear accident has proven that it is crucial to investigate potential severe accidents and corresponding mitigation measures for the sfp of a nuclear power plant (npp) [2, 3]. the necessary mitigation measures, the effects of recovering the sfp cooling system and makeup water in sfp as the accident progresses have also been investigated respectively based on the events of pool water boiling and spent fuel uncovery [3–5]. to use risk insights in the decision-making processes in an adequate manner, it is very important to establish a systematic approach that integrates in a sound, transparent and justifiable manner all the elements required as stated in iaea tecdoc 1436 [6]. requirement 21 paragraph 4.71 of the iaea general safety requirements part 4; states that the regulatory body should carry out a separate independent verification to satisfy itself that the safety assessment is acceptable and to determine whether it provides an adequate demonstration of whether the legal and regulatory requirements are being met [7]. this work looks at all internal possible initiating events that may cause severe accidents in the koeberg nuclear power plant sfp. however, external events such as seismic activity and aircraft crash were not considered for this study. a list of abbreviations is provided in appendix a. 2. probabistic analysis the systems analysis programs for integrated reliability evaluations (saphire) code was used as the modelling tool. this work was carried out by (1) identifying initiating events and scenarios, (2) accident sequence modelling, (3) quantification of top events and (4) analysis of results. the main modelling assumptions for the sfp psa model are listed below: • the sfp water is at initial level of 19.65m. • the initial pool water temperature is at 50◦c. the sfp cooling loop is designed to provide a 50◦c bulk sfp temperature with the maximum component cooling system (rri) temperature and a maximum heat load of 2.88 mw during normal and outage operation. • the model is done for the normal operating mode which is 94.2 % of an 18-month fuel cycle. • no gate seal failure during sfp boiling. • make up from the reactor cavity and sfp cooling system is not considered because of limited capacity. • the demineralised water distribution system (sed), fire fighting water supply (jpp) and the mobile fire fighting system (jps) are the considered make up systems. make up from the reactor cavity and sfp cooling system (ptr) is not considered because of limited capacity 231 https://doi.org/10.14311/app.2022.36.0231 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en sibongiseni thabethe acta polytechnica ctu proceedings figure 1. loss of sfp cooling event tree. figure 2. event tree model of the loss of inventory due to pipe rupture. 3. results and discussions event tree models were developed for the identified initiating events; this was done to better understand accident progression in the sfp resulting from the initiating events. the loss of sfp cooling, loss of sfp inventory due to pipe rupture, loss of sfp inventory due to flow diversion, loss of offsite power and station blackout are the initiating events that were modelled in this study. these initiating events either lead to a loss of sfp cooling, a loss of sfp inventory or both. the success criteria for a loss of cooling or inventory is generally to recover ptr cooling before boiling or before a loss of ptr suction. it also entails making up inventory before ptr suction uncovery and making up inventory prior to fuel uncovery. 3.1. event sequence modelling the loss of sfp cooling initiating event results from a failure of the reactor cavity and spent fuel pool cooling system (ptr) whereby all trains are unavailable for sfp cooling. the loss of sfp cooling event tree is shown in figure 1. following a loss of sfp cooling it will be ideal to recover cooling before boiling occurs (ptrcrecbbnmom). although boiling occurs when the pool heats up to a temperature of 100◦c, it is crucial to maintain the pool level at a safe level, in this case 15,5m. the pumps take suction from the fuel pool at an inlet located below the pool water level (∼ 15.5 m), transfer the pool water through a heat exchanger and return it back into the pool through an outlet typically located below the cooling system inlet and some large distance from it. recovering ptr cooling before the loss of suction at the 17 m level would be ideal in this case (ptrrecbs). if make-up cannot be established and ptr cooling is not restored the sfp level will drop from the 17 m level to the point where it is assumed that fuel damage occurs i.e. at the 9,85m level. the success criterion will be to establish sed make up before fuel damage (sed-mu-bd). in the event where sed is unavailable for make up to the sfp, an operator will be instructed to use the jpp system (jpp-mu-bd). as a last resort, if make up cannot be established through the sed and jpp systems, an operator will have to start the mobile fire pump (jps-mu-bd). this end state indicates that fuel damage has not occurred, the pool is boiling as illustrated by the boil in the event trees. the ok end state indicates that the accident has been successfully mitigated. this state comprises of all states where spent fuel damage has not occurred and either cooling has been re-established or make up to the sfp has been established. the loss of sfp inventory due to a pipe rupture, this initiating event considers the loss of sfp inventory due to pipe breaks in the ptr system. after a pipe rupture event the operator action required is to isolate the break before the loss of the ptr pump suction (ptr-iso-rp). a similar action is required in the case of loss of inventory due to flow diversion, with the operator being required to isolate the ptr flow diversion before the loss of suction (ptrfdiso). the loss of sfp inventory due to a flow diversion, this event is defined as a loss of sfp inventory due to diversion through an interfacing system valve failing 232 vol. 36/2022 koeberg spent fuel pool safety figure 3. event tree model of the loss of inventory due to flow diversion. figure 4. event tree model of losp. figure 5. event tree model of sbo. open or having been left open during ptr operation. the event tree models for the loss of inventory due to pipe rupture and flow diversion are shown in figure 2 and 3 respectively. at the 15.5 m level, ptr pump suction is lost. this means ptr cannot be recovered and make-up via sed, jpp or jps is the only remaining options. this event considers maintaining the sfp water level so that ptr repairs can still be undertaken even after boiling has started. success for this function event is defined as the establishment of sed make-up to the sfp before sfp boiling (sed-mu-bb). should the sed be unavailable for make up to the sfp, an operator will be instructed to use the jpp system (jpp-mu-bb). as a last resort, if make up cannot be established through the sed and jpp systems, an operator will have to start the mobile fire pump (jps-mu-bb). the loss of offsite power is defined as the failure of off-site power supplies. both units are tripped and the emergency diesel generators (edgs) and acacia power station supply are started. off-site power restoration is attempted as soon as possible. edgs will automatically actuate and supply power to the emergency ac buses (lha-b_e). if the edgs start and run, and the operator starts a ptr pump, sfp cooling can be re-established (ptr_clbb19nmom). the ptr pumps must not be restarted with a level of less than 17.0 m in the sfp. the suction of the ptr pumps located at the 15.5 m level in the sfp. below this level ptr cooling cannot be restarted (ptr_c_bs17nmom), therefore make-up is the only remaining option (sed-mu-bd, jpp-mu-bd and jps-mu-bd). the sbo initiating event is the failure of the offsite power supplies, including emergency off-site power from acacia. both units are tripped, the edgs fail to start or run for 24 hours and offsite sources are not recovered. following a sbo, it is desirable to recover power before sfp boiling. this can be done in one of two ways: either the recovery of the grid or by aligning 233 sibongiseni thabethe acta polytechnica ctu proceedings initiating event frequencies initiator time factor conditional initiating frequency (iec) initiating event frequency (ief) loss of sfp cooling (nm) 9,42e-01 9,52e-04 8,97e-04 loss of sfp inventory due to flow diversion (nm) 9,42e-01 3,57e-02 3,36e-02 loss of sfp inventory due to ptr pipe rupture 9,42e-01 1,37e-08 1,22e-08 loss of offsite power (nm) 9,42e-01 4,50e-01 4,24e-01 table 1. summary of initiating event frequencies. figure 6. failure model of the sed system. figure 7. failure model of the jpp system. figure 8. failure model of the jps system. the fifth diesel to either lha or lhb (9lhs_sbo). it is also to recover the grid before sfp boiling 234 vol. 36/2022 koeberg spent fuel pool safety (grid_bb_nm_om) and to start ptr cooling before boiling (ptr_sbo_nm_om). the loss of inventory due to boiling causes the sfp level to drop from the 19.3m level to the 17m level in 37.5 hours in normal mode. success for this function event is defined as the recovery of the grid before the 17 m level is reached (grid_bs_nm_om). as already mentioned, the ptr pumps must not be restarted at a level less than 17m, ptr cooling must be started before loss of suction (ptr_c_sbo_bs_nm_om). it is also desirable to recover the grid before fuel damage (grid_bd_nm_om). the available make up systems before fuel damage occurs are the sed, jpp and jps systems (sed-mu-bd, jpp-mu-bd and jps-mu-bd). the worst-case scenario is a drop in the pool water level to an extent where the fuel assemblies will be uncovered, which is assumed to be at the 9.85m level. this will lead to the fuel being damaged as indicated by the end state fd in the event trees. fuel damage will occur when all possible mitigating actions fail to be implemented. several thermal hydraulic studies on the initiating events have also been performed. these studies are often used to support psa because they reveal crucial information on sfp behaviour [8, 9]. 3.2. accident sequence quantification to quantify accident sequences, fault tree models were used to quantify the function events in the event trees. accident sequence quantification involves firstly the quantification of the initiating event frequencies. the conditional (iec) and initiating event frequencies (ief) were calculated for normal mode of operation (17 months) i.e., a time factor of 94.2 % of an 18-month fuel cycle was applied to the iec to calculate the ief. the iec and the calculated ief were obtained from koeberg nuclear power plant documentation and are shown in table 1. the failure models of the make-up systems take into account human failure events, those function events which require human action were quantified following step by step standardized plant analysis risk-human reliability analysis (spar-h) methodology [10] for quantifying human failure events (hfe). a number of performance shaping factors (psf) are taken in to consideration. these factors include available time, stress/stressor, complexity, experience/training, etc. the failure model of the sed is modelled on the basis that the operator fails to start or align sed make up or the sed isolation valves fail or there is an overall loss of the sed. the loss of the sed results from two events, if the tank leaks externally or the pumps fail to start and run. the failure of the pumps was not modelled in detail and were treated as undeveloped transfers, in this case they are treated as basic events. spar models are generally not as detailed as those contained in licensee pras. the failure probabilities for valves, pumps and tanks were obtained from koeberg internal documentation. the failure model of the jpp system modelled jpp system fails to start and run from standby or operator fails to start or align jpp make-up. the failure model of jps because of the failure of the mobile fire pump from standby or mobile fire pumps suction or discharge valves fail to open or operator ails to establish jps make up. the failure models of the jpp and jps are shown in figure 7 and 8 respectively. the other function events in the losp and the sbo event trees are operator actions such as "operator starts ptr cooling after losp before sfp boiling at 19.3m" (ptr_clbb19nmom) are operator actions and are currently being re-quantified using saphire hra calculator. 4. current and future work the purpose of this work is to do an independent verification of the current koeberg nuclear power plant spent fuel pool psa model. fuel damage is the event consequence state of interest and is considered for quantification. system and failure probabilities were inputted into the model for quantification. since the failure models of the make-up systems involve human or operator action, the function events which require human, or operator action will be re-quantified following a step-by-step standardized plant analysis risk-human reliability analysis (spar-h) methodology for quantifying human failure events (hfe). 5. conclusions this study entails the independent verification of the current koeberg nuclear power plant sfp psa model. initiating events that lead to boiling and potential fuel uncovery were reviewed and investigated. the event tree models for the loss of sfp cooling, loss of sfp inventory, loss of offsite power, station blackout were developed. the accident sequence modelling highlighted the importance of recovering sfp cooling and making up sfp inventory to prevent fuel uncovery. operator or human action is a key component in always maintaining critical pool levels using the available make up systems. through appropriate mitigation measures the increase in the fuel temperature could be stopped, limiting the release of fission products to the environment and preventing severe accident that could result in significant damage to the spent fuel assemblies in sfp. references [1] f. niehaus. use of probabilistic safety assessment (psa) for nuclear installations. safety science 40(1-4):153-76, 2002. https://doi.org/10.1016/s0925-7535(01)00047-9. [2] g. mignot, s. paranjape, d. paladino, et al. large scale experiments simulating hydrogen distribution in a spent fuel pool building during a hypothetical fuel uncovery accident scenario. nuclear engineering and technology 48(4):881-92, 2016. https://doi.org/10.1016/j.net.2016.06.005. 235 https://doi.org/10.1016/s0925-7535(01)00047-9 https://doi.org/10.1016/j.net.2016.06.005 sibongiseni thabethe acta polytechnica ctu proceedings [3] x. wu, w. li, y. zhang, et al. analysis of the loss of pool cooling accident in a pwr spent fuel pool with maap5. annals of nuclear energy 72:198-213, 2014. https://doi.org/10.1016/j.anucene.2014.05.030. [4] s. carlos, f. sanchez-saez, s. martorell. use of trace best estimate code to analyze spent fuel storage pools safety. progress in nuclear energy 77:224-38, 2014. https://doi.org/10.1016/j.pnucene.2014.07.008. [5] x. wu, w. li, y. zhang, et al. analysis of accidental loss of pool coolant due to leakage in a pwr sfp. annals of nuclear energy 77:65-73, 2015. https://doi.org/10.1016/j.anucene.2014.11.010. [6] international atomic energy agency, risk informed regulation of nuclear facilities: overview of the current status, tecdoc series (iaea-tecdoc-1436), 2005. https://www-pub.iaea. org/mtcd/publications/pdf/te_1436_web.pdf. [7] international atomic energy agency, safety assessment for facilities and activities, general safety requirements, no. gsr part 4 (rev. 1), 2016. https://www-pub.iaea.org/mtcd/publications/pdf/ pub1714web-7976998.pdf. [8] k.-i. ahn, j.-u. shin, w.-t. kim. severe accident analysis of plant-specific spent fuel pool to support a sfp risk and accident management. annals of nuclear energy 89:70-83, 2016. https://doi.org/10.1016/j.anucene.2015.11.024. [9] j. c. de la rosa blul, p. mcminn, a. grah. analysis of the inherent response of nuclear spent fuel pools. annals of nuclear energy 124:295-326, 2019. https://doi.org/10.1016/j.anucene.2018.10.014. [10] w. j. galyean, a. m. whaley, d. l. kelly, et al. spar-h step by step guidance, technical report inl/ext-10-18533, rev. 2, idaho national lab. (inl), idaho falls us, 2011. bautechnik 98(7):475-81, 2020. https://doi.org/10.2172/1027888. 236 https://doi.org/10.1016/j.anucene.2014.05.030 https://doi.org/10.1016/j.pnucene.2014.07.008 https://doi.org/10.1016/j.anucene.2014.11.010 https://www-pub.iaea.org/mtcd/publications/pdf/te_1436_web.pdf https://www-pub.iaea.org/mtcd/publications/pdf/te_1436_web.pdf https://www-pub.iaea.org/mtcd/publications/pdf/pub1714web-7976998.pdf https://www-pub.iaea.org/mtcd/publications/pdf/pub1714web-7976998.pdf https://doi.org/10.1016/j.anucene.2015.11.024 https://doi.org/10.1016/j.anucene.2018.10.014 https://doi.org/10.2172/1027888 acta polytechnica ctu proceedings 36:231–236, 2022 1 introduction 2 probabistic analysis 3 results and discussions 3.1 event sequence modelling 3.2 accident sequence quantification 4 current and future work 5 conclusions references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.35.0027 acta polytechnica ctu proceedings 35:27–36, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the railway 4.0 concept – the vision of a digital railway in the czech republic (cr) martin leso czech technical university in prague, faculty of transportation sciences, konviktská 20, 110 00, prague, czech republic correspondence: lesomart@fd.cvut.cz abstract. the paper presents a new alternative approach to the modernization of control and safety technologies on the railway network in the czech republic, referred to as the "the railway 4.0 concept". the presented concept is based on the comprehensive introduction of digital technologies that fulfill the vision of intelligent transport systems on railways (its-r). the paper describes the architecture and basic components of the system based on the distributed technologies with the centralization of logical elements. the paper also places the given issue in the context of current developments in the implementation of the etcs system in the czech republic. keywords: etcs, frmcs, gsm-r, intelligent transportation systems on railway, railway digitalization. 1. introduction the long-planned and long-term implementation of the ertms/etcs system is currently underway, not only on the czech railways. this technology brings, in particular, increased safety in controlling the movement of rail vehicles, i.e. the elimination of error on the part of the train driver. however, the ertms/etcs cannot be seen only as a safety enhancement system, or as a way of meeting one of eu’s interoperable requirements – a requirement to implement a single train protection system. the ertms/etcs technology can clearly be seen as a fundamental element enabling change in the approach to more efficient utilization of the railway system and its implementation must not only follow the safety parameter but must also lead to increased capacity and usability of the railway by the czech transport system. therefore, the question today is not whether to introduce such a system, but how to introduce these new technologies so that they bring real benefit to the railway function and overall transport in the czech republic. 2. history of the railway 4.0 concept the railway 4.0. concept was created as a technological elaboration of the vision "intelligent transport systems on railways" (its-r) 1, which was defined by the sdt working group. the vision of its-r 2 defined the fundamental vision of railway technology 1the its-r vision was added into the strategic materials of the ministry of transportation in 2016 within the implementation plan of the action plan for the development of intelligent transport systems (its) in the czech republic by 2020 (with an outlook to 2050) as project no. 63, zd-20, pilot verification of its-r technology 2http://www.sdt.cz/dokumenty/2015_pozicni_dokument_ sdt_rozvoj_its_na_zeleznici.pdf development. it should clearly lead to the concept of intelligent vehicles, i.e. vehicles which, thanks to the on-board technologies etcs, ato, gsm-r/frmcs, 4g/5g, will enable the full safety of their operation, including automation and optimization of their driving. at the same time, they will enable their full involvement in logistic transportation systems in both freight and passenger transport. finally, they will also provide services to passengers in the form of voice and data communication. however, the infrastructure aspect of these technologies, which must supply adequate technologies in the form of, in particular, radio networks and an appropriate level of etcs train control technology based on at least application level 2 with the possibility of using application level 3, must adapt to this concept of intelligent vehicles. at present, it is not a question of whether this vision of intelligent transport systems will be realized. current global trends, based on the present level of technology and the generally increasing demands for higher railway usability, clearly confirm this direction. similar projects addressing the issue of comprehensive digitalization and modernization of railways can be found in most of the eu’s neighboring countries. the digitales stellwerk 3 project in germany can be mentioned, which is a part of a comprehensive plan for digital railways 4. the traditional swiss railways sbb 5 are dealing with a similar program in the smartrail4.0 project. the comprehensive etcs and gsm-r deployment program is already being implemented by the norwegian railways bane nor 6. the main issue concerning the modernization of 3https://de.wikipedia.org/wiki/digitales_stellwerk 4https://digitale-schiene-deutschland.de/en 5https://www.smartrail40.ch/ 6https://www.banenor.no/prosjekter/prosjekter/ ertms/ 27 https://doi.org/10.14311/app.2022.35.0027 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en http://www.sdt.cz/dokumenty/2015_pozicni_dokument_sdt_rozvoj_its_na_zeleznici.pdf http://www.sdt.cz/dokumenty/2015_pozicni_dokument_sdt_rozvoj_its_na_zeleznici.pdf https://de.wikipedia.org/wiki/digitales_stellwerk https://digitale-schiene-deutschland.de/en https://www.smartrail40.ch/ https://www.banenor.no/prosjekter/prosjekter/ertms/ https://www.banenor.no/prosjekter/prosjekter/ertms/ martin leso acta polytechnica ctu proceedings the railways in the czech republic should therefore be a discussion on how to manage this necessary generational and comprehensive change of technologies in a sufficiently short time, with reasonable financial resources, so that we can take advantage of the benefits that these technologies undoubtedly bring as soon as possible. the aim and the evaluation criteria of modernization strategies should therefore be an increase in the usability of the railway in terms of the speed requirements, infrastructure capacity as well as the accuracy and availability of railways, enabling effective involvement in integrated transport systems throughout the czech republic. in this context of a comprehensive approach, the safety parameter that is most closely linked to the etcs system needs to be seen as a necessary and indispensable condition that is a natural part of the technologies being implemented. in this paper, the 1st of the areas defined in the itsr vision is discussed and further elaborated – control and command technology of railway transport. this technology can be considered the most essential for the function of rail transport. without its systematic introduction, full-fledged railway digitization can no longer be considered. the proposed approach define basic technological elements that should be use by all digital railway applications. subsequent applications or supporting tools forming the overall "digital world" of the railway will then be addressed on this technological infrastructure. however, the issue of related other worlds defined in the its-r vision and its integration into the overall architecture of the "digital world" of the railway is already outside the scope of this paper, so it will not be discussed further here. the railway 4.0. concept was created while the author of this article was active in the safety commission of the ministry of transport of czech republic (mdcr), which was established by the minister of transport of the czech republic doc. ing. karel havlíček, ph.d. mba in 7/2020, in response to a tragic accident at pernink railway station. within the framework of negotiations, a mdcr-o130 vision of the implementation of the etcs and gsm-r systems to all railway lines and vehicles by 2040 was introduced. from the outlined material 7 it is clear that by 2040 all railway lines in the czech republic should be equipped with etcs and gsm-r systems. this political demand, emphasizing in particular the requirement to increase the safety of the entire railway network of the czech republic, was the subject of a solution proposed to address this issue by experts from the správa železnic, státní organizace (hereinafter "sz"). the proposal 8 envisages possible solutions through comprehensive modernization of some major railway lines (most also intended for electrification) via etcs technology at application level 2, on less busy lines it 7https://www.mdcr.cz/media/media-a-tiskovezpravy/bezpecnostni-komise-hlavni-i-lokalni-trati-jiz-zn 8the proposal for the implementation of the etcs system in the czech republic is presented today by the document "plan of modern safety of the czech railways" proposes implementation by repairing existing safety devices and supplementing them with etcs at application level l1, and for the least busy regional lines it proposes supplementing with simplified variants etcs l1 ls and etcs l1 stop. in the case of the etcs l1 version, the gsm-r line radio is also not introduced, on the contrary, the functional use of the trs analogue radio system, which is installed due to the repair of existing srv radios, is practically expanded. however, these simplified variants of the solution only respond to the requirement for accelerated implementation of a safety function preventing trains from an unauthorized passing through a prohibiting signal. nevertheless, they cannot be considered as technologies that would make the railway more usable, including new technologies, as defined by the its-r concept. the railway 4.0. concept brings an alternative proposal for the possibility of a comprehensive approach to the introduction of digital technologies fulfilling the its-r vision on the entire railway network in the czech republic, respecting both the requirements of increasing railway safety and the requirements for higher usability and connectivity of railways for the czech transport system. this paper provides a basic description of the architecture and key components of this concept. when dealing with the complex concept of technology alteration, it can be assumed that the network-wide introduction of these technologies can be managed in a sufficiently short time and at a reasonable economic cost. however, the prerequisite is a sufficiently timeless conceptual planning for the replacement of current oldies technologies, and the subsequent rapid and systematic introduction of digital technologies into real practice. 3. aspects of introducing modern technologies to the czech railway when deploying new technologies based on etcs and gsm-r/frmcs systems, it should be had in mind that the implementation of etcs, which in principle deals exclusively with train safety, as a train safety device having both track-side and mobile parts, is directly dependent on the technology of infrastructure safety devices. these devices are not part of etcs and are therefore do not fall under interoperable specifications and must be addressed specifically by each infrastructure manager. these technologies can thus constitute a significant part of the infrastructure investment that etcs technologies can build on. due to the nature of the etcs system, there are application conditions (functional and safety) that determine which infrastructure safety equipment it is possible or appropriate to apply the etcs system to. when determining the concept of modernization, it is therefore essential to first address the question of the initial situation in the czech republic. unfortunately, 28 vol. 35/2022 the railway 4.0 concept even representatives of the sz in the mdcr safety commission did not provide detailed information on the current parameters of infrastructure technologies. therefore, it is necessary to describe this topic rather in the context of the current approach to the modernization of these technologies. from the current publicly available information and experience of the author of this article, it can be stated that in the czech republic there have traditionally been massive investments into control and safety technologies, especially on corridor networks of approx. 1500 km, which are the backbone railway system included in the ten-t network. nevertheless, there are systems in this part of the infrastructure dating back to the turn of the 19th and 20th centuries, specifically electromechanical safety devices. many railway stations and sections also have installations of relay technology, a track-side safety device, which have been installed since the 1950s. however, the implementation of etcs, application level 2, which is preferred on corridor networks, requires the replacement of these old technologies, as the deployment of etcs would be too complicated and in some aspects also unfeasible for safety reasons. the connection of the etcs system to the newer technologies of electronic interlocks installed at the turn of the millennium can also be complicated due to the moral outdating and unpreparedness of the technology to be connected to the etcs system. specifically, these are, for example, electronic and hybrid interlocking cores, which require data communication with the radio block centre (rbc) and sufficient computing power. as these technologies currently have more than 20 years of active life and are technically and morally outdated, they are currently being replaced by new electronic systems that allow connection to the etcs system. it is evident that the modernization of infrastructure and safety equipment is not easy, it is difficult to solve in the range of corridor lines, and it is very time-consuming and financially demanding. in the czech republic there are currently about 9700 km of tracks in operation. so far, no systematic and comprehensive change of safety technologies has taken place on a larger segment of the off-corridor lines. many of them have so far been modernized due to a lack of funds from their generated profits. nevertheless, a number of national and regional lines, on which most of the necessary comprehensive modernization of the infrastructure has also taken place, have new electronic safety equipment installed. some electronic safety devices are also installed as part of the repairs of outdated safety devices. however, it can be stated that most off-corridor lines (national and especially regional) in the czech republic are burdened by considerable historical underinvestment in control and safety technologies, where mechanical and electromechanical equipment from the turn of the 19th and 20th centuries, as well as relay technologies, are located. 1700 km of regional lines are not equipped with any safety equipment and operation on them is controlled according to the regulation for simplified management of rail transport sz d3 with a minimum of infrastructure technologies. it is therefore a fundamental strategic question how to implement new technologies in a sufficiently short time and as efficiently as possible, which will enable the extension of new etcs and gsm-r technologies to the entire railway network in the czech republic. the key question is certainly also whether it "pays off" to install relatively expensive technologies on lowcapacity regional lines, with current traffic up to tens of trains per day, or whether simpler and cheaper technologies should be applied instead. in this context, it is important to realize that the main role of railways should be primarily the implementation of backbone transport, which forms a unique alternative transport system to the road network, providing both passenger and heavy goods transport. the construction and operation of this railway infrastructure brings relatively high costs both for the infrastructure manager as well as for the carriers of rail vehicles. it is therefore necessary to consider the issue of efficient use of this infrastructure and to consider whether the absence of safety equipment significantly (unduly) restricts the utilization of this infrastructure, and the installation of simplified safety technologies will not significantly increase its utility in terms of capacity (supply offer), as well as its speed. all this can devalue investments in infrastructure and rail vehicles. consequently, this will inevitably lead to increased investment and external costs for higher use of the road network, so savings from the cancellation or reduction of rail transport will certainly not bring overall savings from the point of view of the state or regions. meanwhile, more intensive use of railway transport is currently clearly defined in long-term policies both in the czech republic and in the eu. the inhabitants of the czech republic are currently seeking housing outside large cities for socio-economic and environmental reasons, and it is clear that rail transport can ensure a sustainable way of connecting rural and urban agglomerations. in the area of freight transport, an increase in demand for local freight services in connection with the collection of municipal waste or individual package shipments can also be expected in the coming years. therefore, it is necessary to perceive investments in railways in the context of the transport needs of the whole czech republic, not separately as a solution only to the issue of safety or the introduction of an interoperable etcs system. the introduction of new technologies on regional lines can bring the potential to increase speed and usability even on such underused lines. however, it is an important prerequisite that transport on such lines be implemented as an overall transport concept, where the regional railways will form a fast capacity backbone of the region’s transport system. a specific problem in the czech republic in this context may be regional lines, which to an increased 29 martin leso acta polytechnica ctu proceedings extent face many crossings in a density of up to approx. 1 crossing per 1 km. in addition to the safety problem (road-rail collision), crossings often also limit the speed of rail vehicles, thus limiting the usability of the railways. within the mdcr safety commission, the task of a comprehensive solution for safety at crossings was also assigned. in practice, this means that approx. 150 crossings per year were identified for the implementation of a crossing safety device, typically a separate island crossing safety device. the installation of such crossing facilities is relatively expensive, amounting to around 10 million czk per installation, and in most cases does not lead to an increase in line speeds, due to the existence of other crossings with insufficient safety among other things. from the above description of the initial state of current technologies and the current method of alteration, it is clear that the introduction of new etcs and gsm-r/frmcs technologies in the current railway environment in the czech republic is an extremely complex problem and is a major technical and economic challenge, which must be promptly implemented in the coming years. the railway 4.0. concept brings an alternative comprehensive approach to solving this issue, reflecting the current state of technology in the czech republic based on current modern trends in the field of safety devices and communications. 4. a description of the railway 4.0 concept the railway 4.0. concept proposes a possible technological solution, which is based on the current knowledge of digital technology and reflects developments in this area abroad. the architecture of railway 4.0. is shown in figure 1. the architecture can be divided into the following groups of technologies, which form integral elements of the system: • line transmission and power supply networks – they consist mainly of optical communication and metallic power cables, which form the backbone infrastructure • voice and data communication – enables data and voice communication with on-board technologies, staff, and passengers on board the train • object controllers – a universal programmable element, ensuring the execution of input/output commands controlled by an integrated safety device • integral safety device – centralized logic of all control and safety functions, principally integrating the function of pzz, tzz, szz, rbc • superior operational and dispatch control of traffic – connection to systems enabling remote traffic control and technology management presented architecture of the railway 4.0 concept. cannot be considered final. it is necessary to assume that the number of elements and their applications will continue to develop dynamically according to current system requirements. the technologies discussed in this paper are the basic principal parts necessary for the implementation of the issue of control and command of railway traffic, which is crucial for ensuring rail transport. however, the basic technology defined in the concept (optical line transmission and power supply networks, communications and object controllers) forms the basic infrastructure technology for the implementation of all related digital technologies. it is therefore appropriate that the construction of this basic infrastructure respects the necessary synergies of all technologies, thus enabling the sharing of basic infrastructure elements. as part of the full digitalization of the railways, it will be a matter of supplementing a number of diagnostic and telemetry systems in the infrastructure in order to ensure high availability and safety of the entire railway system. 4.1. line transmission and supply networks the creation of the fundamental infrastructure enabling the transmission of control and diagnostic information and the supply of electricity to technologies is a key part of the implementation of digital technologies. railway is a line system and it requires solution to these technologies along its whole length otherwise it would not be able to ensure the system functionality as defined in the its-r vision. therefore, there are vast and thus high-cost technologies from the point of view of investment and operation. the digital technology being currently implemented requires quality line networks. it is therefore necessary to project these networks with maximum synergy with infrastructural technologies to use investment costs as effectively as possible. when making a long-term conceptual planning of technology modernization at all railways in the czech republic we can assume that the current cable routes are – or in the nearer future will have to be – fixed with optical cables around all railways because the existing railway cable network in most cases consists of outdated communication cables. due to their construction they are mostly at the end of their lifecycle and it is not worth repairing them. currently the standard is that along railways optical transmission networks are laid optical cables. optical routes are thanks to their properties (price, quality of data transmission) an essential part of solution to safety systems and communications. therefore, the railway 4.0 concept is based on building these backbone data routes along all railways in the form of multifibre optical cables. optical data connections serve for connecting object controllers (ocx), bts-r system gsm-r, bts-p of public operators or for further telematics technologies (cameras, detectors of obstacles occupancy, radio beacons etc. . . ) which for their 30 vol. 35/2022 the railway 4.0 concept operation require to gain/provide the railway system with corresponding data. the railway 4.0 concept proposes distributed technologies and therefore at each spot of external technology (crossing, switch, signal, points machine, axel counter point device) the relevant number of optical fibers will be ducted for connecting these components through an object controller. transmission elements (industrial switches) system will be a part of the line transmission network, through which a reliable transmission layer both in inter-station and route sector will be implemented. optical fibers can be also used for other functions mainly of diagnostics and telemetric feature. the application that is currently being discussed and within pilot projects verified is diagnosis of infrastructure faults (typically broken tracks or other defects) or detection of rolling stock failures (typically flat wheels). optical fibers can be also used as elements for diagnosis, monitoring and telemetric of rock or trees falls on superstructure or possibly as a protective perimetric system for protecting railway buildings from unauthorized enter. as a further part of the line routes the railway 4.0 concept suggests supplementing these routes with a power supply line as well. this should serve for supplying all exterior elements and technologies. those are particularly object controllers and peripheral devices of signaling installations controlled by them. it is also necessary to provide electric supply for sufficient support of technologies of radio transmission systems bts-r or public operators bts-p. the current solution to the supply of external technologies lies mainly in connecting public electric distribution systems. for this purpose, long power lines, which have an independent supply line, have to be built, consumption sub-measurement has to be set up. due to the reliability of electric power supply each object has to be provided with a back-up accumulator with corresponding technology of chargers and air condition and heating if necessary. the railway 4.0 concept comes with the possibility of line supply by a route cable 9 that is ducted in excavation together with optical routes within the entire inter-station sector. compared to the solutions so far, the railway 4.0 concept foresees the use of 400v dc supply. this supply technology is currently wide-spread in the systems of photovoltaic power plants and electric vehicles. the prerequisite of such an attitude is the use of low wattage technologies, particularly object controllers and their control peripheral devices. as low wattage technologies we can consider also the independent bts both gsm-r and public operator networks that without control and transmission peripheral devices do not exceed the wattage of 150w to 300w. the object controllers themselves and transmission elements 9such a solution is known from older corridor technologies where the route supply system was made of the line 6kv/75hz dealing with both tract circuits supply and decentralized safety technologies as for example automatic block or crossings. however, this was very expensive and complex solution. in most cases also do not exceed the wattage of 300w. the dimension of such supply networks has to be laid down in detail on the basis of the number of particular supplied objects. according to this, the number and placement of supply and back-up powers of ups is set. the choice of this technology seems to be favorable because of the following reasons: • the supply voltage 400v dc does not require special cables, the technology of electric distribution systems, back-up and protections of similar nowa-days built buildings can be used, for example photovoltaic systems or electric vehicles. • supply back-up can be favorably implemented at railway station buildings where in most cases there is not any problem with connection to a public power line. • the use of dc fitting enables connection of power boosting supply powers anywhere along the supply route which increases acceptable wattage of technologies. from the point of view of costs for acquisition and operation of line networks it is appropriate to build such routes not only for the purpose of technological railway systems but also these networks to enable sharing cable routes for example for the internet within the area or coverage of the area by public operators. current optical cables provide a large number of optical fibers for reasonable prices. compared to the cost of earth-moving – excavation and cable lying (the price of labor-intensive services and machinery) the price of optical cable is negligible. it is, therefore, desirable to lay a sufficiently large number of cables and optical fibers. railway is attractive for lying optical routes for commercial purposes of area coverage with data services. particularly regional lines can enable cheaper and easier implementation of data cables in the so-called white places of the czech republic than lying optical cables in the landscape including the necessity of negotiating with many landowners. for a railway the investments in building and running such networks may become – under the condition the legislative and economic rules are unambiguously set – potentially profitable. 4.2. voice and data communication the railway 4.0 concept introduces the principle of a complex coverage of all lines with data and voice transmissions of mobile radio as a basic means for safe railway traffic control. for control and securing the railway the gsm-r system is currently being built, which in the forthcoming years (appr. after 2030) will have been gradually replaced by the frmcs system. the railway 4.0 concept assumes that the gsm-r system (and frmcs in the future) will be carried out by the system of distributed technologies which is based on the principle shown in figure 2. in an appropriate object of the infrastructure, for example at 31 martin leso acta polytechnica ctu proceedings figure 1. the architecture of railway 4.0. 32 vol. 35/2022 the railway 4.0 concept figure 2. an example westermo data switch. a station, a unit – the so-called digital module (dm) – will be installed that via an optical cable enables to control other up to 6 remote transmitters, the socalled radio remote head (rrh) which create the base station bts-r. the maximum distance you can control with one dm module is up to 30 km. in most cases of terminal regional lines this distance will be fully sufficient though from a practical standpoint at many lines it will be necessary to increase the number of rrh and dh modules due to expected worse radio signal coverage of the area. the supply of rrh would be solely from the line supply cable. as for providing sufficient back-up we can consider elimination of local battery elements which would make the fitting simpler and cheaper. this solution enables minimalization of technology in infrastructure comparing the current state of the gsm-r technology at the corridor network where each bts contained all transmitting elements, accumulators of a power supply and air-conditioning. this coverage of tracks by the gsm-r system will be integrated in a standard way into the already built gsm-r network whose capacity is sufficient to cover all tracks in the czech republic. in the case of the system gsm-r (frmcs) installation we can also – in a synergic way – use the technologies (line transmission and data networks, columns, switch boards) for coverage a track and its adjacent area by public operators’ signal. it is assumed that this technology will become a distributed one that will not require local expensive and demanding technologies. figure 3. architecture of distributed technology gsm-r, source: kontron transportation s.r.o. 4.3. object controllers an object controller ensures the function of input/output interface toward outdoor technologies (crossing scanner, railway barrier, signal, point machine, etc.) on the basis of data communication with integral safety device. in principle it is the so-called smart periphery that carries out the transmission of data reports onto physical inputs/outputs. the object controller can also contain basic logical and diagnostic functions. for connection of specific technologies (for example control of the lights of level crossing signals pzz, axel counter, electric motor of a railway barrier or points machine) in the object controller specific control interface can be fitted that is adjusted to the requirements of external technologies. from the standpoint of a railway safety device this is technology that has to meet the requirements for electronic safety device placed outdoors that meets the requirements for safety of electronic systems with hw and sw solution on sil=4 level. the technologies of the object controllers themselves have to meet the conditions of outdoor environment as they will be placed in an outdoor closet by a track. it is appropriate to design an object controller as a device that is universal and independent of a concrete product. generally, for such devices we use sign cots (commercial off-the shelf) or mots (modifiable off-the-shelf). those are products that are offered to more producers as commercial, mass-produced ones. in contemporary technological development dealing particularly with automation and robotization of industry, also known as industry 4.0, the market with such products has been developing fast and we can assume that it will also include producers that make products meeting railway requirements. the communication interface between oc and izz should be user-open (the owner/operator should get entire product documentation package). this should ensure the independence of oc technology suppliers. currently 33 martin leso acta polytechnica ctu proceedings for example the european initiative eulynx 10 is dealing with such general specification. this initiative associates 13 managers of infrastructure in order to prepare standardization of interfaces and elements of signaling systems. figure 4. an example himatrix controller hw that can be used as object controller. 4.4. integral safety system (izz) the integral safety system is used to implement all logical functions necessary for control and safety of railway operation. current technologies require the installation of classical safety devices, such as pzz – level crossing safety devices, tzz – line safety devices, szz – station safety devices in local objects (usually relay rooms in station buildings), where all the technology is concentrated and connected by cabling to external elements. the railways 4.0. concept is based on distributed object controller technology with the centralization of logical functions into one center. it thus works with a so-called integral safety system that concentrates all logical functions. it is, therefore, a universal computing hw equipped with specific software integrating pzz, szz, tzz functions. it must also integrate the functions of the etcs rbc system at application level l2 or l3. as the izz will use etcs technology of application level l2 or l3, enabling the transition to direct vehicle control via gsm-r (in contrast to the current method of vehicle control via signals, which also uses etcs at application level l1), we can apply more advanced approaches to train control in rbc, where trains are allowed to leave in direct "successions" with dynamically assigned travel authorisation according to real train movement conditions, without having to wait for the release of an entire intermediate block section or a long track section. this can be achieved either by equipping the infrastructure with a larger number of axle counters 11 or by introducing the so-called "floating block" rbc function within the etcs application level l3. the contribution of the etcs l3 application level is that 10https://www.eulynx.eu/ 11typically every 200m, especially at a track junction before entering a station it will allow this train control function to be implemented with a minimum of infrastructure detection features, and also more efficiently in the context of ’flexible traffic control’, without the constraints of infrastructure detection technologies. etcs application level l3 requires an additional train integrity check 12, which can bring investment and operational savings for the infrastructure at the cost of increasing the equipment especially of freight trains 13. however, it is required that exclusive operation with etcsequipped trains and continuous gsm-r / frmcs track coverage are in place. figure 5. an example himatrix controller hw that can be used for izz functions. more efficient traffic control can also be achieved by massive introduction of the ato system 14, which will enable further optimization of train speeds with the positive effect of maximum infrastructure utilization, stabilization of operations and significant energy savings. in order to increase capacity, especially on single-track lines, it is possible, using the abovementioned technologies and the principle of traffic control, to set up approx. 2km long passing tracks with two tracks, e.g. in the intermediate block section, which can very effectively ensure the so-called "fast passing" of trains. these modern control approaches, 12the requirement to check the integrity of the train is addressed in tsi subset026-3v360 in chapter 3.6.5.2 and can be provided by technical means on the train (a unit train is indivisible or the train is equipped with end of train detection technology) or by the operating staff. 13end of train systems are implemented around the world under the name "end of train device" (eod) or end of train (eot) and are massively operated for example on the canadian pacific railway system in the usa. in the eu, this system has not yet been uniformly specified. 14ato – automatic train operation – a system of automatic train guidance on the optimum travel trajectory. the system is currently ready for approval as part of the new tsi subset. 34 https://www.eulynx.eu/ vol. 35/2022 the railway 4.0 concept conditioned by the introduction of etcs technologies at application level l2 or l3 together with gsm-r / frmcs, allow a significant increase in infrastructure capacity and train speed on both corridor lines and single track lines compared to the current situation. the izz concept should allow for easy sw configuration of the system so that in case of a change in the sw configuration (typically the addition of another controlled area or modification of an existing one), only the relevant part of the sw configuration is changed and subsequently tested without affecting the remaining functional part of the izz. according to findings from abroad, izz can also be considered as a so-called "cloud" solution, providing very high computing and data space and very high availability and reliability of the solution. it goes without saying that the requirements for this part of the safety system will also be at sil=4 level for hw and sw functions; and it is absolutely necessary to ensure adequate protection of the entire computer solution against cyber-attacks. the railway 4.0 concept assumes use of the applicable interoperability specifications and the relevant "subset" for ertms / etcs system, with which the technologies used must be in full compliance. therefore, standardly implemented certified technologies such as gsm-r components, etcs mobile parts, balises and rbc functions are assumed. the discussed technology, especially concerning transmission and power supply elements, object controllers or the technology of integrated security devices, is not directly controlled by interoperable requirements. therefore, compliance with the target requirements for safety and reliability, which are controlled mainly according to the requirements of the set of cenelec standards of the en 5012x series, and compliance with the target requirements of safety and reliability management for the given system are expected. these detailed parameters are not the subject of this paper, and can be analyzed in the phase of specifying the requirements for the relevant technological units. 4.5. master operational and dispatching traffic control master operational and dispatching technologies are used to control traffic on a given track or in an area where distributed technology will be deployed according to the railways 4.0. concept. it is assumed that the current approaches and technologies that are built and operated today within the central dispatching departments will be used. however, given the assumption that etcs application level l2 or l3 technology will also be used, it can be assumed that this technology will also be adapted to these new rail traffic control options. it can also be assumed that master technologies for automatic train operation (ato systems) will also be used and should be integrated into this part of the operational traffic control system. significant developments in this area can thus also be expected. 5. practical implementation of technologies according to the railways 4.0 concept when implementing distributed technologies with centralization of logical functions according to the rail 4.0. concept, the construction must start with the building of the backbone line infrastructure, including both optical and power cables. when building this network, the location of the outdoor elements of the tracks, where optical and power lines will be routed to object controllers, and the location of bts radio networks should already be known. this basic infrastructure must, therefore, be built robustly enough and with sufficient consideration to ensure that any further modifications to the superstructure will not disrupt it and will enable it to continue to function flawlessly even if the infrastructure is upgraded over time. the next step is to build local object controllers. their construction is not too demanding as it is an industrial technology that can be installed in industrial distribution boards for outdoor installation. poles for the installation of bts for gsm-r and gsm public operators will also be built. a great advantage of the implementation of distributed technologies can be seen especially in solving the problem of level crossing safety, which is a significant source of danger and slowdown of traffic on regional lines. the railways 4.0. concept practically allows every level crossing to be equipped with an adequate level of external features that will ensure crossing safety pursuant to legal requirements. the benefit of this solution, apart from the increase in safety, is above all the possibility of running trains at the maximum line speed without any restrictions resulting from the currently insufficient level of level crossing safety along the entire track section. at the same time, the investment and operating costs of building such level crossing safety system will locally be significantly lower, because the costs of most of the functions implemented locally at the level crossing (level crossing logic, power supply and data communication, etc.) are already solved by means of line communication and power supply networks and centralization of logic functions in the izz. last but not least, a backup power supply must also be built. buildings that already have or allow for the new installation of a power connection from the public power grid will be given priority. station buildings generally meet this requirement without the need for extensive internal occupation of the building as is necessary in the case of conventional installation of station safety equipment. when using a 400v dc supply voltage distribution system, it is also possible to reinforce the lines in other buildings on the track if necessary; however, calculations to date show that at 35 martin leso acta polytechnica ctu proceedings standard station distances of up to approx. 7km, it will be sufficient to supply this track power line only bilaterally from adjacent station buildings. with a sufficient high-quality implementation of this network, a very long service life of at least 50 years or more can be expected. when connecting object controllers to external peripherals, a unified type method will be used, which will be specified for each type of external peripheral separately. the final activation of the technology is realized only by configuring the izz, where specific inputs/outputs of object controllers are assigned to relevant sw objects, which are further handled by the izz software. this is expected to greatly simplify and speed up not only the design of these technologies but also the final wiring and testing of the entire system. it is evident that with regard to the size of the railway network which must be covered by the etcs and gsm-r / frmcs system in relatively short time, the construction and expansion of this technology cannot be combined with complex modernization activities on the railway construction part, which progress at a significantly slower pace (for technological and economic reasons) than that of line cabling and bts radio networks construction. on the contrary, the approach based on the railways 4.0. concept significantly accelerates the coverage with modern control and safety technologies of even those parts of the infrastructure where no heavy medium repair work or possibly complex modernization of the infrastructure has been carried out so far. a system with distributed technologies has the conceptual requirements to be significantly more adaptive and modifiable in its ability to implement configuration changes or renewal of individual parts than is usual in the case of static locally centralized installations. this is because only partial segments of the technology are replaced when a change or renewal is implemented, while the costliest segments related to power, data networks and central logic remain unchanged or undergo partial configuration changes. 6. economic parameters of the railways 4.0 concept the costs for the implementation of the railways 4.0. concept were processed by the author of the article in the form of an estimate of investment costs according to the sfdi rules 15 for the purposes of the discussion of this proposal in 2020 in the safety commission of the ministry of transport and communications working group infrastructure. this estimate based on a model example of a typical regional line, suggests that the cost of the presented comprehensive view of modernizing railway control and safety is not significantly higher than the currently planned costs, even in case of application of significantly simplified 15https://www.sfdi.cz/pravidla-metodiky-a-ceniky/ cenove-databaze/, sborník pro oceňování železničních staveb ve stupni studie proveditelnosti a záměr projektu etcs l1 ls or stop installations without gsm-r track-side radio. in the case of the railways 4.0. concept, the scope of technology, functions and usability of the railway is comprehensively addressed at the highest possible level of current knowledge, with the possibility of also meeting future requirements, e.g. autonomous operation of trains without drivers. the costs of track safety may range from 5 to 7 million czk/km 16 of single-track line and thus respect the estimated amount of financial resources defined in the vision of the ministry of transport o130 for the introduction of the etcs system until 2040. in addition, lower operating costs can be expected when this technology is implemented. this is due both to the use of highly reliable technologies with detailed diagnostics and lower operating costs ensuring a reliable and secure power supply and centralized logical parts of the system. 7. conclusion the presented railways 4.0. concept proposes a new alternative approach to the modernization of control and safety technologies on the railway network of the czech republic, which should enable a comprehensive introduction of digital technologies on the railway network in the czech republic. the railways 4.0. concept is technological elaboration of the its-r vision, while it not only pursues the issue of modernization of safety systems, introduction of the etcs system enabling the increase of the level of safety of railway operation in the czech republic, but also sets the technological environment enabling the increase of the usability of the railway and it readiness for greater inclusion in the integrated transport systems of the regions in the czech republic. the concept was discussed at the level of the mdcr and the sz in 9/2020. based on the decision of the director general of sž bc. jiří svoboda, mba, sz took over the official patronage of this railways 4.0. concept and will further develop this solution within the framework of joint work group railways 4.0. comprising experts from sž, ctu in prague, faculty of transportation. the work group is currently developing this concept to the level of implementation project documentation and is working towards the subsequent implementation of a pilot verification of this concept on the real infrastructure of the sž test track. 16approximately 200 k euros to 270 k euros per km of installation 36 https://www.sfdi.cz/pravidla-metodiky-a-ceniky/cenove-databaze/ https://www.sfdi.cz/pravidla-metodiky-a-ceniky/cenove-databaze/ acta polytechnica ctu proceedings 35:27–36, 2022 1 introduction 2 history of the railway 4.0 concept 3 aspects of introducing modern technologies to the czech railway 4 a description of the railway 4.0 concept 4.1 line transmission and supply networks 4.2 voice and data communication 4.3 object controllers 4.4 integral safety system (izz) 4.5 master operational and dispatching traffic control 5 practical implementation of technologies according to the railways 4.0 concept 6 economic parameters of the railways 4.0 concept 7 conclusion acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0382 acta polytechnica ctu proceedings 38:382–388, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague a case study of implementing a renovation concept matthias haase zurich university of applied sciences, institute of facility management, grüental / ra, p.o. box, ch-8820 wädenswil, switzerland correspondence: matthias.haase@zhaw.ch abstract. there is an untapped potential for reducing ghg emissions by district renovation. it needs a detailed energy master planning (emp) of the district and support for the decision-making processes. if applied in a comprehensive manner it can contribute significantly to reducing energy consumption and thus to a long-term sustainable development of districts and climate neutrality in our cities. the emp typically includes combinations of energy supply and consumption, but it is equally important to understand the different solutions for district renovation. the multi-owner structure in many districts requires another set of solution finding that is embedded in potential analysis, stakeholder analysis, participative planning, and multi-actor-management. a district near winterthur, switzerland was used as a case study. site visits and structured interviews with key stakeholders were used to collect data which was then used to model the technical-economic situation and to determine the possibilities for the future. keywords: district renovation, renovation packages, building renovation, energy supply. 1. introduction previous research shows, that renovation strategies on building level need to be derived from the energyefficient renovation of buildings and the use of renewable energy sources to reduce the energy supply of regions or cities. the combination of energy efficiency and renewable energy sources covers both energy supply and demand in the built environment. in this sense, the renovation of buildings is a suitable strategy to reduce demand, while the use of renewable energy aims to reduce carbon dioxide emissions from the energy supply system. however, identifying technical solutions is not enough to implement large-scale renovation strategies and to achieve the predicted decarbonization of the building stock. the renovation rate in europe is clearly below the 3 percent annual target [1, 2]. some of the main barriers to remediation relate to remediation costs and access to finance, as well as complexity, awareness, stakeholder management and supply chain fragmentation [1, 3]. as a result, business models are relevant to implementation and acceleration of renovations. seddon et al. [4] define “business model” as the outline of essential details of a firm’s value proposition for its various stakeholders, and the activity system the firm uses to create and deliver this value proposition [5]. in other words, a business model is the abstraction of a strategy, focused on the system of activities through which a firm creates economic value. opportunities to reduce greenhouse gas emissions through neighborhood renovations are largely untapped. it requires not only in-depth energy planning (emp), but also the support of decision-making processes [4]. this can not only significantly contribute to reducing energy consumption and ensuring the location of energy infrastructure (production, distribution, storage), but also to long-term sustainable development of districts and climate neutrality of our cities. 1.1. energy and emissions to be able to reduce ghg emissions in the built environment it is important to focus on the reduction of co2 emissions from the operation of the buildings [6]. the reduction of energy use should come from the implementation of efficiency measures through energy renovation of the building stock. another possibility is the decarbonisation of the energy supply with on-site renewable energy measures. however, in renovation planning it is often unclear which energy supply options are available and what influence the energy renovation has on co2 emission reductions. 1.2. reduction paths and its elements a two-phase approach is proposed to reduce carbon dioxide emissions: first, reducing energy consumption with efficiency measures in connection with the renovation of the building stock. secondly, co2 emission reductions from on-site renewable energy supply. when it comes to costs and financing, it is crucial to relate different measures with different stakeholders. energy supply is a political (municipal) matter, but the renovation of residential buildings mostly depends on the owners [6]. in order to achieve carbon reduction goals, it is important to find ways to involve 382 https://doi.org/10.14311/app.2022.38.0382 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 a case study of implementing a renovation concept figure 1. the settlement seen from above with eight building blocks containing 51 row houses. homeowners in long-term carbon reduction investment strategies [7]. this case study was modelled with different power supply options and the performance was simulated [1]. the performance gives energy consumption over a year and with energy prices the costs were calculated. the cost savings could then be used in part to finance an energy-efficient renovation. in the base case, fossil fuels (oil) are delivered to the district and different power supply options offer energy and cost savings compared to the base case. after that, the energy saving measures are calculated, which lead to energy cost savings for the building owner. finally, local renewable energy sources are integrated into the roofs of the buildings which provide electricity production and income when sold to the utility company. the final investment costs are calculated and presented to the building owners. 2. case study 2.1. situation settlement 51 in dinhard was built in a first stage in 1974 and in a second stage in 1977 as a cooperative settlement. eight blocks contain 51 row houses as shown in figure 1. the row houses have different sizes and different number of rooms (3.5 rooms, 5.5 rooms and 6.5 rooms) resulting in different row house types. some of the houses are already being retrofitted with new windows and additional roof insulation, but so far no district renovation (of all houses) took place. the heating center in front of block f next to the swimming pool (in white in the middle lower part) supplies all row houses with heating and hot water. the oil boiler with 465 kw heat output was replaced in 2019. 2.2. energy consumption an average of 99 370 liters of oil (average for the years 2000 to 2014) was consumed in the district. with an average energy reference area of a row house of 140 m2, an energy figure for heat ew = 140 kwh/(m2a) can be found. this value is somewhat below the swiss average for older, non-refurbished residential buildings (ew avg = 160 kwh/(m2a)) [3]. 2.2.1. energy supply option three different building standards (standard 1, 2 and 3) with different peak power was modelled. for each building standard, the following energy supply options were considered [1]. • oil boiler (status quo), centralized system with distribution system. • oil boiler for heating, hp system for dhw. ▷ oil boiler for heating, centralized system with distribution system. ▷ decentralized heat pump system for dhw. • ground source heat pump, centralized system with distribution system. • pellet boiler, centralized system with distribution system. 2.2.2. energy, costs and emissions the simulations provided energy consumption results for each option. these were taken as input for cost and emission calculations. the following assumptions were considered as shown in table 1 (energy costs and emissions), table 2 (investment costs), table 3 (maintenance costs) and table 4 (energy costs). 3. results the results of the case study are presented in energy use, investment and maintenance costs as well as energy costs, and ghg emissions. the link between costs and emission reduction will then be presented to the investors/housing owners. 383 matthias haase acta polytechnica ctu proceedings energy carrier costs emissions electricity 0.15 chf/kwh 0.25 kg co2eq/kwh pellets 375 chf/t 0.01 kg co2eq/kwh oil 1.05 chf/l 0.28 kg co2eq/kwh table 1. energy costs and emission factors of different energy carriers. unit chf oil boiler central oil boiler gshp pellet boiler+ dec. hp boiler electricity 0 7500 34290 0 pellets 0 0 0 60000 oil 84000 68250 0 0 el. for pumps and motors 2000 1800 2000 2000 total energy 86000 77550 36290 62000 table 2. investment costs of different supply options (in chf). unit chf oil boiler central oil boiler gshp pellet boiler+ dec. hp boiler cleaning, control, fees, services 2500 2000 1000 2500 repair fond 3400 6570 5000 6480 total 5900 8570 6000 8980 table 3. maintenance costs (in chf). unit chf oil boiler central oil boiler gshp pellet boiler+ dec. hp boiler electricity 0 7500 34290 0 pellets 0 0 0 60000 oil 84000 68250 0 0 el. for pumps and motors 2000 1800 2000 2000 total energy 86000 77550 36290 62000 table 4. energy costs (in chf). annual energy use type 1 type 2 type 3 type 4 [kwh] block a 135564 5 1 block b 135564 6 block c 170680 1 5 1 block d 194972 1 7 block e 146388 1 5 1 block f 189878 1 6 1 block g 124643 0 5 block h 97168 0 4 sum 1194858 3 1 43 4 table 5. energy use. unit kwh/(m2a) type 1 type 2 type 3 type 4 heating 118 118 118 118 dhw 22 22 23 24 electricity 24.4 31.6 36.3 39.12 sum 164.4 171.6 177.3 181.12 table 6. energy use in different housing units. 384 vol. 38/2022 a case study of implementing a renovation concept renovation measure description investment costs [chf] roof tiles + wind barrier + 30 cm insulation + vapourbarrier 653 158 facade + new windows + insulation panels + doors 705 944 ventilation ventilation unit + heat exchanger + ductwork 324 602 pv 1339 pv modules + fastening structure + inverter +cabling 709 322 pvt 1339 pv modules + inverter + cabling, partly integrated with 354 m2 solar thermal collector modules solar thermal collector (5, 6, 7, 8 m2 for building types 1, 2, 3, 4 respectively) + piping + storage tank 852 422 table 7. renovation measures with investment costs. operational energy energy savings co2 emissions co2 emissions savings [kwh] [kwh] [kg co2] [t co2] ref 1212971 322688,35 reno 1 roof 938563 274409 242573,05 80,1 reno 2 roof+fassade 801354 411617 202069,10 120,6 reno 3 roof + facade + balanced ventilation 664146 548826 161565,16 161,1 reno 4 reno1+pv 821331 391641 224425,58 127,8 reno 5 reno1+pvt 740929 472042 201384,15 150,8 reno 6 reno2+pv 679185 533786 183157,37 169,1 reno 7 reno2+pvt 603721 609251 160880,20 191,3 reno 8 reno3+pv 541977 670995 142653,43 209,6 reno 9 reno3+pvt 466512 746459 120376,25 231,8 table 8. renovation options. 3.1. energy there are a total 8 different blocks with four different types of row houses (type 1, 2, 3, 4). energy use in the eight different building blocks is illustrated in table 5. a total of 1195 mwh is used. table 6 summarizes the energy use per m2 heated floor area for heating, domestic hot water (dhw) and electricity. the different types have different annual energy use between 172 kwh/(m2a) (type 1) and 181 kwh/(m2a). table 7 shows the different renovation measures for the eight buildings. the first measure focuses on the roof and adds a vapour barrier, 30 cm insulation, a wind barrier and new tiles. the investment costs for this measure were calculated to 653 158 chf. the second measure focuses on the facade and includes changing doors and windows as well as insulation panels underneath the windows. the investment costs for this measure were calculated to 705 944 chf. the third measure includes a balanced ventilation system consisting of a ventilation unit per house, heat exchanger and ductwork. the investment costs for this measure were calculated to 324 602 chf. the fourth measure includes a pv system on all roof surfaces, including fastening structure and cabling as well as inverters. the investment costs for this measure were calculated to 709 322 chf. the fifth measure includes a pvt system, adding to the pv system solar thermal collectors (5, 6, 7, 8 m2 for building types 1, 2, 3, 4 respectively), piping and a storage tank. the investment costs for this measure were calculated to 852 422 chf. the five measures were combined to nine different renovation options (reno 1 to reno 9). it follows step by step renovating the roof, the façade, and then integrating the different technologies pv and pvt in different options as listed in table 8. the energy use and savings were calculated for heating, dhw and electricity separately. in the case of pv, a separate calculation included a share of self-consumption (30 %), and the economic implications which arise from different tariffs for purchasing and selling electricity [8]. 385 matthias haase acta polytechnica ctu proceedings figure 2. economic evaluation of the renovation options (pbp and npv). in the two columns on the right the co2 emissions of each renovation option is shown together with the potential savings compared to the reference case. 3.2. economic evaluation figure 2 shows the net present value (npv) and payback period (pbp) of the different renovation options (reno 1 to reno 9) without the different options for energy supply. it can be seen that the pbp varies between 16 (reno 4) and 25 (reno 3) years. with the chosen discount rate of 3.5 % and a lifetime of 25 years two renovation options have positive npv (option 4 and 5) and are economically feasible. however, renovation option (reno 1, reno 7, reno 8 and reno 9) are very close to a positive npv and in combination with a wood pellets boiler they become also economically feasible. the payback periods are lowest for renovation options (reno 4) (15.5 years) and (reno 5) (15.4 years). the highest pbp shows renovation option reno 3 (24,5 years). eight renovation options show a pbp below 20 years (reno 1, reno 4, reno 5, reno 6, reno 7, reno 8 and reno 9). 3.3. ghg emissions table 8 gives the ghg emissions and the savings of the different renovation options. we included investment as well as maintenance and energy costs of the supply options. figure 3 shows investment costs over the saved co2 emissions of the different renovation options for different supply options (see table 3) over the period of 25 years. the lowest costs per t co2 provides the renovation of the roof (reno 1) in combination with a wood pellets boiler (100 chf/t co2). the other renovation options are more expensive with (reno 6) providing the highest costs (489.4 chf/t co2). further, costs are always highest for the oil boiler 386 vol. 38/2022 a case study of implementing a renovation concept figure 3. economic evaluation of the renovation options with different supply options. option, while the two gshp solutions rank second and the wood pellets solution is always the cheapest. the renovation options (reno 2), (reno 4) and (reno 5) (options 4 and 5 have a positive npv) with wood pellets boiler provide the second lowest costs (205, 196 and 214 chf/t co2 respectively). 4. conclusions the results of the case study clearly show the potential to save energy and greenhouse gas emissions. correlation with the necessary investment and maintenance costs and energy costs shows that at least two basic improvement options are advantageous. here, the economic advantage of the energy savings of roof renovation is combined with regenerative energy production on the roof. when presenting the emission reductions and avoided costs, it becomes clear that this effect can be strengthened by additional fees for greenhouse gas emissions. in this case, as the fees increase, depending on the co2 level, more renovation options become financially viable. several renovation options are close to economic profitability, and especially the renovation options (reno 1, reno 7, reno 8 and reno 9) become feasible with a combination of wood pellet boilers. this is due to several reasons. first, the co2 payment savings are greater. this is due to the higher co2 savings of the delivered energy. with the co2 coefficient of wood pellets, 0.036 kg/kwh has been calculated, compared with the co2 coefficient of oil of 0.295 kg/kwh. the economic benefit of each saved kwh consists of energy costs and a higher co2 payment saved. finally, the renovation versions reno 4, reno 6 and reno 8 produce electricity. while the self-consumption rate (30 %) lowers energy costs, the excess of electricity is sold to the grid. this lowers energy costs and increases co2 payments. in the renovation options reno 5, reno 7 and reno 9, additional hot water is produced in pvt collectors. the share of own consumption (43 %) is used directly for hot water in buildings. the remaining part (57 %) is not used. however, it can be used for other purposes (such as heating the swimming pool or heating other neighbours). we want to explore the possibilities of using this extra heat for other purposes in the next phase of research. in this work, each building has a small part of the roof with a solar heat collector (5, 6, 7, 8 m2 for building types 1, 2, 3, 4). if all roof areas were supplied with pvt, the excess heat could be stored in seasonal storage increasing the use of locally produced renewable energy and the district’s self-sufficiency. acknowledgements this work has been supported by the project “ped – positive energy districts” of the zhaw. expert discussions within the iea ebc annex 75 and annex 83 are highly appreciated. references [1] m. haase. implementation strategies for renovation concepts based on participative planning. iop conference series: earth and environmental science 1078(1):012081, 2022. https://doi.org/10.1088/1755-1315/1078/1/012081 [2] i. artola, k. rademaekers, r. williams, j. yearwood. boosting building renovation: what potential and value for europe?, european union, brussels, 2016. [3] k. laffont-eloire, n. peraudeau, s. petit, et al. sustainable business models for the deep renovation of buildings, stunning – sustainable business models for the deep renovation of buildings, 2019. [4] p. b. seddon, g. p. lewis, p. freeman, g. g. shanks. the case for viewing business models as abstractions of strategy. cais 13:25, 2004. https://doi.org/10.17705/1cais.01325 [5] bpie. europe’s buildings under the microscope, building performance institute europe, brussels, 2011. [2022-06-22]. https://bpie.eu/wp-content/uploads/ 2015/10/hr_eu_b_under_microscope_study.pdf [6] m. haase, d. baer. constraints, stakeholders, and framing goals in energy master planning between neighborhood and district. in a. bisello, d. vettorato, d. ludlow, c. baranzelli (eds.), smart and sustainable planning for cities and regions, pp. 3–13. springer 387 https://doi.org/10.1088/1755-1315/1078/1/012081 https://doi.org/10.17705/1cais.01325 https://bpie.eu/wp-content/uploads/2015/10/hr_eu_b_under_microscope_study.pdf https://bpie.eu/wp-content/uploads/2015/10/hr_eu_b_under_microscope_study.pdf matthias haase acta polytechnica ctu proceedings international publishing, cham, 2021. https://doi.org/10.1007/978-3-030-57764-3_1 [7] buildings performance institute europe (bpie). scaling up deep energy renovations. unleashing the potential through innovation & industrialization, 2016. [2022-06-22]. https://www.bpie.eu/wp-content/ uploads/2016/11/bpie_i24c_deepretrofits.pdf [8] european commission dg energy. financing the energy renovation of buildings with cohesion policy funding, 2014. [2020-04-01]. https://ec.europa.eu/ energy/sites/ener/files/documents/2014_ guidance_energy_renovation_buildings.pdf 388 https://doi.org/10.1007/978-3-030-57764-3_1 https://www.bpie.eu/wp-content/uploads/2016/11/bpie_i24c_deepretrofits.pdf https://www.bpie.eu/wp-content/uploads/2016/11/bpie_i24c_deepretrofits.pdf https://ec.europa.eu/energy/sites/ener/files/documents/2014_guidance_energy_renovation_buildings.pdf https://ec.europa.eu/energy/sites/ener/files/documents/2014_guidance_energy_renovation_buildings.pdf https://ec.europa.eu/energy/sites/ener/files/documents/2014_guidance_energy_renovation_buildings.pdf acta polytechnica ctu proceedings 38:382–388, 2022 1 introduction 1.1 energy and emissions 1.2 reduction paths and its elements 2 case study 2.1 situation 2.2 energy consumption 2.2.1 energy supply option 2.2.2 energy, costs and emissions 3 results 3.1 energy 3.2 economic evaluation 3.3 ghg emissions 4 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0088 acta polytechnica ctu proceedings 40:88–92, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague dependence of mechanical and thermal properties on the composition of lightweight gypsum composites hana sekavováa, b, ∗, zdeněk prošeka, c, pavel tesáreka a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague, czech republic b knauf praha spol. s r.o., mladoboleslavská 949, 197 00 prague 9, czech republic c czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: hana.sekavova@fsv.cvut.cz abstract. one of the ways to use the treated gypsum waste is in gypsum mixtures for the production of gypsum blocks. gypsum blocks can be used in standard interior as well as exterior applications and are made of gypsum, water, pp microfibres and are lightened with foam, which is created from a foaming additive and water. above all, the amount of foam significantly influences the bulk density of the resulting material, and the bulk density then has a major influence on the mechanical and thermal properties. the use of pp microfibres had a positive effect on the overall stability of the foamed structure, which resulted in an increase in compressive strength while maintaining good thermal insulation properties. keywords: mechanical properties, thermal properties, gypsum, lightweight composite. 1. introduction the trend nowadays is to use recycled materials generated in the construction industry and reuse them in the construction industry. in this way, mineral resources are protected and there is no need to extract new non-renewable raw materials. such materials include gypsum, which is up to 100 % recyclable. one of the most common uses of gypsum today is primarily in the manufacture of plasterboard (pb). the issue of recycling of construction pb is related to its contamination with pb of different types from different manufacturers (different compositions) and other materials used in the production of pb structures – dry constructions [1, 2]. these include mineral insulation, metal profiles, pur foams, residues of wipers, etc. another problem is related to the biodegradation of pb [3–5]. the above contamination complicates the return of this material for pb re-manufacturing, which is quite demanding in terms of the quality of the input recycled material, technological processes, etc. therefore, efforts are made to find suitable materials and products based on recycled construction pb that are not so susceptible to changes in the quality of pb recyclate [6, 7]. one way to use recycled pb is in the form of lightweight blocks to be used for building envelopes. blocks for building envelopes today are primarily designed to have a certain balance between mechanical properties (primarily compressive strength, modulus of elasticity, etc.) and thermal performance. the most commonly used thermo-technical parameter is the thermal conductivity coefficient, which can then be used to calculate the thermal resistance and possibly other characteristics that are already part of the building design or an optimal envelope composition. the composition is possibly supplemented by thermal insulation materials. the advantage of gypsum-based materials is the fact that the gypsum can be modified, either by additives and admixtures, to modify the resulting performance of the hardened material. the primary objective in the design of a new gypsum mixture (dry or wet) is to modify the internal structure (in terms of internal arrangement and ratio of mass and excitatory pores), and thus to change the bulk density of the fresh gypsum slurry, i.e. of the hardened and subsequently dried samples, where the mixture/material/product already has the resulting material properties [8–10]. in terms of achieving optimal mechanical properties (compressive strength) and thermal properties (thermal conductivity coefficient), these are completely opposite effects. in terms of compressive strength, we need a material with a minimum number of pores and, conversely, for the thermal conductivity coefficient, we would need a material full of small pores to make the whole structure lighter. at first glance, quite different properties and requirements can be achieved with the optimum ratio of the components used – water, gypsum or gypsum primary inert binder, using a foaming additive and microfibres. in general, chemical additives such as setting retardant, plasticizing additive, etc. can be used. a properly applied foaming additive (in terms of the amount of foaming additive, the 88 https://doi.org/10.14311/app.2023.40.0088 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 40/2023 dependence of mechanical and thermal properties on ... (a). demolition of a plasterboard wall. (b). waste inside the smart container. (c). typical waste after emptying from the container. figure 1. demolition of a plasterboard wall, waste inside the smart container, typical waste after emptying from the container, in this case green. element symbol atomic conc. weight conc. oxide symbol stoich. wt conc. o 66.09 46.39 s 17.09 24.05 ca 16.81 29.56 figure 2. chemical analysis of recyclate by sem eds analysis. amount of water, the quality and stability of the foam) will lighten the gypsum matrix, ideally by small pores that do not clump and are evenly distributed throughout the volume of the mass, and the microfibres will “hold” the entire foamed and lightened structure during the initial stages of setting and hardening and will also improve the resulting mechanical properties. another advantage of the whole concept of producing gypsum blocks based on recycled gypsum waste is that waste from the production of gypsum blocks can be incorporated into the mix as a primarily inert material. a further advantage is the fact that when the gypsum blocks are refurbished or end-of-life, they can be reused or recycled again and used for the production of new blocks [11]. 2. materials and samples the composites were composed of recycled pb materials. the recyclate came from the construction site, was sorted using a “smart” container and contained mainly “ordinary” white (grey) pb from knauf praha, spol. s. r. o. (figure 1). construction pb waste was treated on a mobile recycling line at lavaris s. r. o. in libčice nad vltavou. the treated pb recyclate was delivered to the faculty of civil engineering of the czech technical university in prague in closed plastic cans. at the faculty of civil engineering, analysis of this material was carried out using a pheonom xl scanning electron microscope. the recyclate was examined using a bse detector in combination with eds analysis (figure 2). the results show that the purity of the recyclate is greater than 95 % of caso4. from the point of view of efficient utilization and verification of the technological possibilities of production of lightweight blocks in laboratory conditions, it was advisable that the calcination rate was as high as possible, for this reason a selected part of the imported pb recyclate was further calcined to achieve at least 95 % calcination rate of the total weight of the recyclate samples, based on the weight of the binder, i.e. including any impurities. the calcination was carried out under laboratory conditions using a laboratory oven heated to 120 °c until the mass was stable, corresponding to the evaporation of water or the stoichiometric difference between the calcined 89 h. sekavová, z. prošek, p. tesárek acta polytechnica ctu proceedings samples plaster water w/p fibres type of foam foam addition [g] [g] [-] [g] [g] f 3.0 2.1 0.7 0 none 0 g 3.0 2.1 0.7 15 none 0 h 3.0 1.8 0.6 15 in-situ 60 i 3.0 1.8 0.6 15 in-situ 45 j 3.0 1.8 0.6 0 in-situ 30 k 3.0 1.8 0.6 15 in-situ 30 l 3.0 1.8 0.6 15 in-situ 15 table 1. used mixtures and their composition. gypsum and the gypsum. the degree of calcination was verified on selected samples using so-called drying scales. to ensure the stability of the foamed structure, specially manufactured pp microfibres from trevos were used. microfibres have 32 µm in diameter and 4 mm in length. other properties are following: average tensile strength, ≤3.0 cn/dtex (∼272 mpa); average elongation, ≤50 %; density, 910 kg/m3; young modulus of elasticity, ∼4 gpa. based on previous experiments, several sets of gypsum-based specimens were designed to verify the dependence of water coefficient and foam amount on the final properties of the lightweight gypsum composites. for comparison purposes, reference sets without lightening (f and g) were also created to better identify the degree of lightening and compositional adjustment. other sets were specific in the amount of foaming agent used (15, 30, 45 and 60 g). the foam was produced so-called caused in-situ. some sets additionally contained pp microfibres. for practical reasons, a water coefficient of 0.7 had to be used for sets f and g, and for the other sets it was reduced to 0.6. for each set, 6 cubes with an edge of 100 mm were made. see table 1 for details. 3. experimental methods prior to testing, all the test samples were artificially dried at 40 °c using a hot air oven for 3 days until the weight stabilized. they were then stored in a desiccator at laboratory temperature and 0 % humidity for 24 hours. the samples thus prepared were subsequently tested. first, the samples were tested to determine the bulk density and thermal properties. testing of thermal properties was performed on heat transfer analyzer isomet 2104 (applied precision) equipped with surface probe api210411 and api210412. the investigated thermal property was the thermal conductivity coefficient λ. the surface probes had measuring range from 0.04–0.3 w/mk, and 0.3–2.0 w/mk, and accuracy of ±5 % was used. such a device applies a dynamic method which is based on response monitoring of examined material on heat flow impulses. measurements were made on 100 × 100 × 100 mm samples and each measurement was made on a different side of the cube. each sample was measured three 850 900 950 1000 1050 1100 1150 1200 f g h i j k l b u lk d e n s it y [ k g /m 3 ] figure 3. comparison of the average values (with standard deviations) of the bulk density of the tested gypsum mixtures. times and the resulting value was averaged and the standard deviation determined. the temperature of the samples during the measurement was equal to 22±1 °c (the same as ambient air). finally, destructive tests were performed. the destructive tests were carried out using a heckert fp100 loading frame with displacement-controlled loading at 0.1 mm/min. a uniaxial compression test was carried out on cubic samples. compressive strength was calculated from the maximum force reached during the test, fc,max, as: fc = fc,max ab . (1) 4. experimental results after the solidification and hardening process, a porous system was formed in the hardened gypsum mass, which defines the bulk density. after drying the samples from the water necessary for processing (from a technological point of view), the value of the bulk density of the naturally wet gypsum solid structure was determined. the comparison of the average bulk density values in figure 3 shows the influence of the fibers between sets f and g. the lowest value was measured for sets i and h with the highest amount of foaming additive added, while the lowest value was measured for set k. 90 vol. 40/2023 dependence of mechanical and thermal properties on ... 0 1000 2000 3000 4000 5000 6000 7000 8000 9000 f g h i j k l c o m p re s s iv e s tr e n g th [ k p a ] figure 4. comparison of the average values (with standard deviations) of the compressive strengths of the tested gypsum mixtures. figure 4 compares the average compressive strengths. the highest average values were obtained for the reference samples f and g without foaming, but a value below 7 mpa was also obtained for set l with 15 g of foaming additive and pp fibres. the lowest average values were obtained for set h and i, where their values fell below 3 mpa. figure 5 compares the average thermal conductivity coefficient values for the tested gypsum mixtures. in terms of this parameter, set k with 30 g of foaming additive and pp fibres has the worst values around 0.5 w/mk. the best average value of 0.35 w/mk was measured for sets h and i. for most of the other sets, values around 0.4 w/mk were obtained. 5. conclusions the paper presents the first part of the results in the field of effective use of treated gypsum plasterboard waste treated by a recycling line and its subsequent use in the form of lightweight blocks for building envelopes. the described reuse method seeks to efficiently use the gypsum plasterboard recyclate and transform it into a new building product that can compete with conventional materials, such as the classic aerated concrete blocks that are commonly used today. calcination was also verified in the laboratory using a laboratory dryer at temperatures up to 120 °c. in the next part of the project, the aim will be to add a new calcination unit to the recycling line, which would effectively refine the treated plasterboard recyclate and convert it into a gypsum binder with similar properties to a conventional building plaster. acknowledgements this paper was financially supported by the tacr trend 3 research a project under the number fw03010054 “recycling and transformation of construction plasterboard waste into new products for construc0.000 0.050 0.100 0.150 0.200 0.250 0.300 0.350 0.400 0.450 f g h i j k l c o e ff ic ie n t o f th e rm a l c o n d u c ti v it y [w /m k ] figure 5. comparison of the average values (with standard deviations) of the thermal conductivity coefficient of the tested gypsum mixtures. the values were determined for naturally moist samples in a laboratory environment at a temperature of 25±2 °c and a relative humidity of 50±3 %. tion and value-added applications” and by czech technical university in prague under no. sgs project sgs22/089/ohk1/2t/11 “buidings materials and construction under view of experiment”. the authors also thank knauf prague ltd. for the supplied materials. references [1] m. a. pedreno-rojas, i. flores-colen, j. de brito, c. rodríguez-linán. influence of the heating process on the use of gypsum wastes in plasters: mechanical, thermal and environmental analysis. journal of cleaner production 215:444–457, 2019. https://doi.org/10.1016/j.jclepro.2019.01.053 [2] a. ahmed, k. ugai, t. kamei. investigation of recycled gypsum in conjunction with waste plastic trays for ground improvement. construction and building materials 25(1):208–217, 2011. https: //doi.org/10.1016/j.conbuildmat.2010.06.036 [3] h. sekavová, j. herrmann, z. prošek, et al. the influence of the use 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and materials, vol. 204, pp. 1237–1241. trans tech publ, 2012. https://doi. org/10.4028/www.scientific.net/amm.204-208.1237 [11] a. erbs, a. nagalli, k. q. de carvalho, et al. properties of recycled gypsum from gypsum plasterboards and commercial gypsum throughout recycling cycles. journal of cleaner production 183:1314–1322, 2018. https://doi.org/10.1016/j.jclepro.2018.02.189 92 https://doi.org/10.1016/j.resconrec.2016.01.014 https://doi.org/10.1016/j.jobe.2021.102524 https://doi.org/10.1080/09593330.2015.1102329 https://doi.org/10.1007/s13369-020-04362-4 https://doi.org/10.4028/www.scientific.net/amm.204-208.1237 https://doi.org/10.4028/www.scientific.net/amm.204-208.1237 https://doi.org/10.1016/j.jclepro.2018.02.189 acta polytechnica ctu proceedings 40:88–92, 2023 1 introduction 2 materials and samples 3 experimental methods 4 experimental results 5 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0228 acta polytechnica ctu proceedings 38:228–234, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague hygric performance of new building components for vertical green gardens pavel kopecký czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic correspondence: pavel.kopecky@fsv.cvut.cz abstract. the newly developed box-like panel lightweight element for vertical greenery was designed and built at university centre for energy efficient buildings of the czech technical university in 2020. a test site for testing green facade components is operated since spring 2021. test samples differ in shape, arrangement of functional layers, water distribution means and irrigation patterns. the article presents and discusses measured data at selected test samples during growing season in year 2021. physical quantities, such as surface temperature of plants, moisture content and matric potential of growing substrate, water inflow and outflow, and ambient boundary conditions, were recorded. keywords: vertical green gardens, real-scale experiment, hygric performance, watering patterns. 1. introduction the space next to buildings in urban areas is in particular used for pavements and roads. solid nonporous materials like concrete and asphalt are used. the built environment lacks green plants if compared with rural areas. building enclosures can be seen as potential carriers of green plants. whereas technical solutions for green roofs are established on the market the technical solutions for green vertical walls are still under development. the current state-of-the art of vertical greenery systems is summarized in [1]. vertical green gardens attached to building walls offer some benefits. greenery improves the visual quality of built environment, improves soundproofing properties of building enclosures, protects the building enclosure and the internal environment against external climatic loads and may create new ecological niches. massive usage of green roofs and vertical greeneries within urban areas could mitigate the urban heat island effect [2]. some disadvantages are however associated with vertical green gardens as well. first, it is known to only limited extent which plants are suitable for given type of climate, environmental conditions of site and the system for vertical greenery. a lack of long-term measurement data from real-scale installations exists. second, the raw building materials are needed. for instance, some containers to keep growing substrate at place and supporting structures carrying the greenery are needed. moreover, the irrigation system has to be used in order to compensate for unfavourable environmental conditions at vertical walls. the irrigation system and greenery itself need regular inspections. it can cause non-negligible maintenance costs during operation. it is the subject of current research if some of the disadvantages can be minimized by optimized and reliable technical solutions. this paper deals with newly developed planar boxlike elements for vertical green gardens. various test samples were built and equipped with measurement sensors in 2021. this article presents a side-by-side comparison of measured hygric performance of several test samples in real climatic conditions of the site located in semi-continental climatic conditions. 2. materials and methods 2.1. development of a new technical solution for vertical greenery the goal was to design a technical solution for vertical greenery which will be suited for vertical walls of industrial buildings (e.g. warehouses and production spaces). the walls of modern industrial buildings are currently being constructed as lightweight sandwich structures with low capacity to carry extra loads. therefore, it was required that the proposed solution should be lighter than 50 kg/m2 (i.e. less than half of the system for vertical greenery described in [3]). the requirement of low weight implies the limit on the thickness of soil-type growing substrate (< 5 ÷ 6 cm). the soil-type substrate was believed to be applicable for plants under consideration and for semi-continental climatic conditions. since only one worker should be able to carry the weight of the element, surface area of the element is practically limited to 0.5 m2. water availability is essential for long-term success of plants. the technical solution for vertical greenery must ensure reliable inflow and controlled flow of water through growing substrate leading to uniform and sufficient moisture content. the challenge is to control the vertical redistribution of water due to gravity. the goal is to maintain stable environment with good water availability. the development of a system for controlled distribution of irrigation water in the element was carried out iteratively by trial and error approach. several 228 https://doi.org/10.14311/app.2022.38.0228 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 hygric performance of new building components . . . (a). 26/11/2020. (b). 14/04/2021. figure 1. the first prototypes placed on the test façade. laboratory models of the water distribution system in the characteristic sample were built (06–08/2020). the laboratory experiments tested two conceptual ideas: (1.) water distribution along the upper edge of the element can be provided by a suitable hydrophilic capillary active material, and (2.) water distribution in growing substrate can be enhanced by a drainage layer located at the rear side of the panel (behind soil substrate). testing of watering concepts finally led to an original solution using a block of hydrophilic mineral wool placed along the top edge of the element. two parallel plastic plates with teeth shaped bottom edge are pushed in the cut-outs made in the mineral wool. teeth release drips of water if hydrophilic mineral wool is oversaturated. water drips into the soil substrate which is enclosed in the honeycomb-like structure of interconnected cells made of recycled plastics (cell size 62.5 × 62.5 mm, thickness 4 cm). the growing substrate is firmly hold in cells by jute fabric and metal lattice placed on the external side. water also drips into drainage layer which is located behind growing substrate. the carpet made of recycled polyester was used (thickness 2 cm). water moves down through cells with growing substrate and through carpet by means of gravity. the carpet enables faster water transport than soil substrate. the overflow of water from carpet to growing substrate is created in the middle of the height of the element. the carpet is there discontinued by means of horizontal yet sloped cut. the plastic plate with teeth is placed in the cut and sealed to the rear side of the element. water flow through the carpet is thus diverted from rear side into growing substrate. the first three test samples (scale 1:2, full height, half width) were built during 09/2020. two test samples were seed by grass and hanged on the test façade. the test samples were fully green since the end of 10/2020. the test samples were irrigated weekly in the cold period. grass on both test samples survived cold season in a relatively good shape, albeit showing weaker coverage of the bottom side of both test samples and reduction of leaf area density than in initial state. the fade was mainly associated with one-week long cold period (daily average ambient temperature < 0 °c) at beginning of 03/2021. grass relatively quickly regrown during warmer weeks of spring 2021, even without new seeds (see figure 1b and figure 4a, b). the experience from the cold season showed that capillary redistribution of water in hydrophilic mineral wool does not work reliably. the uniformity of outflow from teeth was dependent on finding the good compromise between total water inflow, size and geometry of mineral wool, and geometry of teeth. moreover, mineral wool suffered from material degradation and changed its hydraulic properties during the cold season. the distribution of water along the top edge of test samples was therefore redesigned during spring 2021. a thicker plastic pipe with low diameter inlet holes (drilled in angle of 120° and regularly distanced to each other) was placed along the top edge of the sample instead the block of mineral wool. the tests have proven that combination of diameters 13 mm (main pipe) and 1.5 mm (individual outlet holes) leads to uniform outflow of water from holes if water pressure exceeds 2 bars. six real-scale test samples were gradually built in 2021. test samples differed in shape, arrangement of functional layers and water inflow distribution devices. five samples were hanged on the test façade of the research centre uceeb. test sample nr. 4 was built and sent to test site of industrial partner. for differences between samples see table 1. main parts of a test sample are depicted in figure 2. test samples nr. 3–6 were built in plastic euro containers (size 80×60×12 cm). containers for samples nr. 1 and nr. 2 were built from scratch using extruded polystyrene and plywood. 229 pavel kopecký acta polytechnica ctu proceedings figure 2. technical drawing of test sample nr. 3. sample dimensions substrate plants inflow overflow 1 12 × 7 cells(76 × 44 cm) s1, 43 mm grass (standard mixture of seeds) block of hydrophilic mineral wool with plastic teeth carpet with overflow (4th row) 2 6 × 15 cells(38 × 95 cm) s1, 43 mm grass + veronica spicata, rudbeckia hirta block of hydrophilic mineral wool with plastic teeth no carpet 3 12 × 8 cells(76 × 50 cm) s1, 43 mm grass + veronica spicata, rudbeckia hirta plastic pipe dn 16 mm with capillaries 2 mm carpet with overflow (5th row) 5 12 × 8 cells(76 × 50 cm) s1, 63 mm grass + arabis alpina, fragaria vesca plastic pipe dn 13 mm with holes 1.5 mm no carpet 6 12 × 8 cells(76 × 50 cm) s2, 63 mm thymus serpyllum, sedum reflexum plastic pipe dn 13 mm with holes 1.5 mm no carpet s1 – 3 vol. grass substrate / 1 vol. acre substrate. s2 – 3 vol. grass substrate / 1 vol. acre substrate / 1 vol. sand, 1 vol. expanded clay table 1. definition of test samples. 2.2. description of test site the test site for green facade components was built at university centre for energy efficient buildings in buštěhrad (uceeb) during spring 2021. the test site is located on south-east oriented window sill of the experimental curtain wall (at the ground level). the space behind test façade is unheated and ventilated by outdoor air. the following sensors were used to monitor conditions in test samples: (1.) temperature sensors (pt1000, accuracy ± 0.15 k), (2.) fdr volumetric moisture content sensors (analog truebner smt 50 and digital smt 100, accuracy ± 3 % vwc, no laboratory tests were performed for calibration prior measurement campaign), (3.) gypsum blocks for measurement of soil water potential (delmhorst gb2, measuring range −0.1 bar to −15 bar), (4.) non-contact temperature sensors (infrared radiometers, apogee si 421-ss). the boundary conditions were measured on the test facade by ems 33h, temperature and relative humidity sensor shielded against radiation, accuracy ± 0.15 k, ± 2 % rh. global solar radiation was measured by ems11, low cost global radiation sensor. global radiation data recorded at roof of the research centre were also available. outflow from test samples 1 and 2 is measured by rain gauge (pronamic pro). positions of sensors are shown in figure 3. 2.3. description of irrigation system the test site was equipped by an automatic irrigation system. the system was used for watering of test samples nr. 1 and nr. 2. the system consists of magnetic valves, balancing valves, water flow meters, compressor, fittings and plastic pipes for water supply to the test facade. the magnetic valves and compressor are both controlled by programmable logic controller (plc). the compressor enables emptying of the system during cold season in order to prevent damage of external pipes and fittings by water freezing. the logic of the controller can be programmed and tested in matlab/simulink environment [4]. the code is then built and run in plc. two standard control schemes were used at first: (a) timer with specification of time and water amount for watering pulse (used in test sample nr. 2) and (b) feedback control with irrigation controlled according to actual measured moisture content. 230 vol. 38/2022 hygric performance of new building components . . . figure 3. position of sensors in test samples (external view on test samples – since 08/2021). feedback control was used in test sample nr. 1. test sample nr. 2 used time controller scheme. morning and evening watering event were usually used in the very warm days. watering event was constituted by two consecutive pulses. the first watering pulse was used to saturate the upper part of the sample. the second watering pulse transported water from the oversaturated top to the bottom part of the sample. test samples nr. 3–6 were irrigated by standard commercially available garden timers. 2.4. a method used for estimation of real evapotranspiration the steady-state thermal balance of the plant surface was used to estimate the real daily evapotranspiration et from a test sample. the following equation can be derived: et = 1 lv ( αsolggt − hre(tae − tre) − ∆t ( hse + 1 r )) [kg/(m2day)], (1) where hse [w/(m2k)] is total heat transfer coefficient at plant surface, hre [w/(m2k)] is radiative heat transfer coefficient at plant surface, r [m2k/w] is thermal resistance between plant surface and ventilated air gap behind a test sample, αsol [-] is absorptivity of plants for short-wave radiation, ggt [w/m2] is daily mean global solar irradiance on the test facade, tre [°c] is daily mean radiant temperature of surrounding surfaces, lv [j/kg] is latent hat of water evaporation (2.5 × 106 j/kg), and ∆t [k] is the difference between daily mean plant surface temperature tpe and daily mean ambient air temperature tae (∆t = tpe − tae). for practical calculation, it is assumed that mean radiant temperature is approximately equal to apparent sky temperature (tre ≈ tsky ). the apparent sky temperature is estimated according to [5]. real evapotranspiration can be alternatively estimated by method [6] as: et = kcet0 [kg/(m2day)], (2) where et0 is reference evapotranspiration, i.e. evapotranspiration from grass surface having fixed crop height, surface resistance and albedo, and kc is real crop coefficient (time variable). 3. results 3.1. visual investigation of test samples the test façade was visited almost every week during warm season 2021. a diary with impressions and comments was written. photographs were regularly taken by a mobile phone and by outdoor camera mounted on the stand in front of the façade. surface patterns recognizable on test samples shows the condition of plants and to some extent moisture content of soil substrate. figure 4 shows selected days in time period 06–09/2021. the visual appearance of test samples in june was promising with nice green colour and no brown coloured spots (see figure 4a). only the half-size test samples built in 2020 shows lower leaf area density at the bottom part but this is believed to be the consequence of spring freezing week. the figure photographed in july already shows some signs of fade at some locations of test samples (see figure 4b). there is a brown spot on the right side of test sample nr. 1 and brown area located at the bottom edge of test sample nr. 3. one month later almost all grass of test sample nr. 1 died (see figure 4c). in september, test sample nr. 1 shows improvement at the bottom and top edge whereas middle part remains brown. brown spots are evident at test samples nr. 2 and nr. 3 in august and september 2021 (see figure 4c and d). 231 pavel kopecký acta polytechnica ctu proceedings (a). 14/06/2021. (b). 19/07/2021. (c). 10/08/2021. (d). 07/09/2021. figure 4. photographs of test façade – selected days in growing season 2021. (a). 04/06/2021. (b). 12/08/2021. figure 5. surface temperature of plants tpe compared with ambient air temperature tae in two selected clear sky days. figure 6. difference of daily mean plant surface temperature and ambient air temperature (6–9/2021). 232 vol. 38/2022 hygric performance of new building components . . . figure 7. evapotranspiration of test samples compared with reference evapotranspiration. visual observation also showed signs that soil substrate contained excessive amount of water, i.e. brightness of the surface, extrusion of water when soil surface was pressed by fingers, muddy fingers after wiping, and frequent water leaks at the bottom edge of test samples. mould growth was observed on the jute fabric of test sample nr. 1. the difference of plant patterns between test samples nr. 3 and nr. 5 deserves comments (see figure 4c and d). seeds germinated more quickly on the upper (and moister) part of sample nr. 5, whereas grass grew better at the bottom (and not so moist) part of test sample nr. 5 (see figure 4, compare c and d). the upper part of test sample is moister since no overflow element was used in test sample nr. 5. test sample nr. 3 seemingly shows the opposite to test sample nr. 5. seeds germinated more quickly on the bottom part whereas grass grew well along the upper edge. in this case, the moister part is located at the bottom. the wetness of the bottom part of test sample was caused by the drainage carpet and overflow element used in test sample nr. 3. 3.2. the surface temperature of plants the measured surface temperature of plants was compared with ambient air temperature in two similar clear sky warm days of growing season (04/07, 12/08). the surface temperature of plants was lower than ambient air temperature by about 4 °c for both test samples in 04/07 (see figure 5a). the surface temperature of test sample nr. 1 exceeded ambient air temperature by 7 °c in 12/08 whereas test sample nr. 3 maintained undercooling of the plant surface (see figure 5b). the evolution of the difference between daily mean plant temperature and daily mean ambient air temperature is depicted in figure 6. 3.3. daily evapotranspiration daily evapotranspiration calculated by equation 1 is compared with reference evapotranspiration et0 in figure 7. 4. discussion the locations with poor plants were documented by photographs (rise of characteristic brown spots visible in photographs). dying of plants occurred during the second half of growing season and should not be attributed with the lack of available water in root area in this case. in fact, the reason for dying of plants was probably opposite – long term excessively high moisture content of growing substrate. the excess water in root area caused the lack of oxygen in growing substrate. when one would like to improve the whole concept, operation of irrigation, technical concept and selection of plants would have to be better balanced for given environmental conditions. the task is difficult due to many design variables coming into consideration, e.g. the geometry of cells, position of protruding holes amongst cells, hydraulic properties, thickness and composition of soil substrate, distance between the upper distribution pipe and the bottom edge of the greenery element, position and presence of the overflow elements, amount of water and watering patterns. 5. conclusions this study dealt with hygric performance of the newly developed building component for vertical greenery. investigations were based on measured data and regular visual observations of real-scale test samples in warm season of year 2021. the initial promising growth of plants was replaced by their gradual decline in the second half of growing season. overwatering is considered to be the main reason for decline of plants. the presented technical solution of the panel element should be therefore further optimized to obtain better environmental conditions for plants. the main design issue is to find trade-off setting between water storage and water transport properties of the element, i.e. to balance the composition of growing substrate, the shape of plastic cells, position of overflow elements and daily watering patterns. last but not least it is evident that better understanding of plants would be helpful. systematic overwatering showed that watering needs of plants should be better predicted. therefore, an algorithm for prediction of daily watering needs was programmed in [4] and embedded in plc. weather forecast apis [7, 8] provide access to data for environmental boundary conditions on site in the upcoming day or days. the first test runs were performed at the end of year 2021. the algorithm will be further tested in year 2022. the test period was limited to growing period of year 2021. the performance of test samples in nongrowing period must be studied carefully in conti233 pavel kopecký acta polytechnica ctu proceedings nental climatic conditions as well. from practical perspective, it is questionable how to protect pipes and fittings against freeze-thaw damage. acknowledgements this work has been supported by project tačr th 03030230. the author is very grateful to radim havlíček for his technical assistance. references [1] g. pérez, j. coma, i. martorell, l. f. cabeza. vertical greenery systems (vgs) for energy saving in buildings: a review. renewable and sustainable energy reviews 39:139–165, 2014. https://doi.org/10.1016/j.rser.2014.07.055 [2] m. santamouris. cooling the cities – a review of reflective and green roof mitigation technologies to fight heat island and improve comfort in urban environments. solar energy 103:682–703, 2014. https://doi.org/10.1016/j.solener.2012.07.003 [3] p. kopecký, j. tywoniak, k. staněk. experimental investigation of green façade components for industrial and storage buildings. iop conference series: earth and environmental science 588(5):052049, 2020. https://doi.org/10.1088/1755-1315/588/5/052049 [4] matlab, 2020. 9.9.0.1718557 (r2020b), natick, massachusetts: the mathworks inc. [5] m. aubinet. longwave sky radiation parametrizations. solar energy 53(2):147–154, 1994. https://doi.org/10.1016/0038-092x(94)90475-8 [6] r. g. allen, l. s. pereira, d. raes, m. smith. crop evapotranspiration – guidelines for computing crop water requirements – fao irrigation and drinage paper 56. [2022-01-20]. https: //www.fao.org/3/x0490e/x0490e00.htm#contents [7] weather forecast api. [2022-01-20]. https://wp2.pvforecast.cz/ [8] weather forecast api. [2022-01-20]. https://openweathermap.org/ 234 https://doi.org/10.1016/j.rser.2014.07.055 https://doi.org/10.1016/j.solener.2012.07.003 https://doi.org/10.1088/1755-1315/588/5/052049 https://doi.org/10.1016/0038-092x(94)90475-8 https://www.fao.org/3/x0490e/x0490e00.htm#contents https://www.fao.org/3/x0490e/x0490e00.htm#contents https://wp2.pvforecast.cz/ https://openweathermap.org/ acta polytechnica ctu proceedings 38:228–234, 2022 1 introduction 2 materials and methods 2.1 development of a new technical solution for vertical greenery 2.2 description of test site 2.3 description of irrigation system 2.4 a method used for estimation of real evapotranspiration 3 results 3.1 visual investigation of test samples 3.2 the surface temperature of plants 3.3 daily evapotranspiration 4 discussion 5 conclusions acknowledgements references my title https://doi.org/10.14311/app.2022.36.0198 acta polytechnica ctu proceedings 36:198–205, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague full probabilistic design of a submerged floating tunnel and format for partial safety factors franka e. m. swaalfa, b, ∗, raphaël d. j. m. steenbergenb, c a delft university of technology, faculty of civil engineering and geosciences, stevinweg 1, 2628 cn delft, the netherlands b the netherlands organization for applied scientific research tno, p.o. box 155, 2600 ad delft, the netherlands c gent university, department of structural engineering, technologiepark-zwijnaarde 60, 9052 gent, belgium ∗ corresponding author: lisa.swaalf@tno.nl abstract. a submerged floating tunnel (sft) can be a promising solution for crossing a deep or wide waterway. this innovative concept however lacks research into its probabilistic design. in this research, the reliability of the tether-stabilized sft is assessed. a first-order reliability method (form) and a monte carlo simulation (mcs) are performed for the limit state functions of the most important failure mechanisms. stochastic variables are chosen so that a target reliability index of 3.8 for a reference period of 50 years is met. the calculated factors from the full probabilistic design are compared with the general recommended partial factors for strength and resistance from eurocode en1990. for the strength mechanisms, the calculated factors are smaller than the factors from eurocode. however, for the equilibrium mechanism, the calculated factor for the unfavorable loading is larger than the factor from eurocode and should be increased by 10% in order to design a safe enough sft. keywords: eurocode en1990, first-order reliability method, submerged floating tunnel, monte carlo simulation, partial safety factors, tether-stabilized, tether slackening, tether yielding. 1. introduction a solution for crossing a deep and wide waterway can be a floating tunnel under the water level: a submerged floating tunnel (sft). this type of tunnel consists of an immersed concrete tube, either attached with anchor cables to the seabed or attached to pontoons floating on the water surface. these two variants can be seen in figure 1. the concept of an sft was patented in the united kingdom in 1886, and the first norwegian patent on the subject was issued in 1923. in the 1960’s a small group of norwegian experts started to evaluate the potentials of the sft concept. more recent, the government of norway made plans for the highway route e39, crossing many fjords, to be ferry-free. this goal could be achieved by constructing sfts crossing the fjords on this route [1]. in this paper, we focus on the reliability of the sft and the applicability of the recommended general partial safety factors from eurocode en1990 [2]. previous research focused mostly on the structural lay-out of the sft. few researchers have addressed the safety of the structure and no probabilistic design has been made before. the experience gained from other civil structures can only partly be used, since sfts differ from common civil structures. in general, sfts have different relevant limit states, different magnitude of consequences and costs involved with failure of relevant limit states, different accuracy of the models that predict the structural response and different load scenarios [3]. the forcing on underwater structures differs from the forcing on structures on land, since the equilibrium of the system is determined by the buoyancy-weight ratio. the general partial safety factors for strength and resistance from eurocode en1990 are calibrated on forces acting on buildings and bridges. it is questionable whether these factors can be applied to the sft [2]. the aim of this paper is to perform a full probabilistic calibration of partial factors for most relevant limit states of the sft. a level iv probabilistic analysis is not performed, since costs are difficult to determine and can vary significantly. accidental loading types are out of scope for the limit state functions, as well as the loads on the land-bored part of the tunnel and the end-joints. this paper is structured as follows: first, the methodology is presented (section 2), in which a typical sft structure is shown, a target β-value is set and the calculation of partial factors is explained. consequently, in section 3, the failure mechanisms are described and all stochastic variables are explained. in section 4, the resulting reliability indices and partial safety factors are discussed, followed by the overall conclusions (section 5) [4]1. 1this paper is based on the authors thesis, reliability analysis of submerged floating tunnels, 2020 [4] 198 https://doi.org/10.14311/app.2022.36.0198 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 36/2022 submerged floating tunnel design for safety figure 1. two variants of an sft figure 2. simple model of sft-element with loading types 2. methodology 2.1. procedure first, a typical sft structure is investigated (section 2.2). in order to assess the reliability of the sft, a decomposition of this system is made (section 2.3). the governing failure mechanisms are determined. consequently, the reliability requirements on element level are set. a target β-value is selected for a certain reference period and limit state functions are formulated (section 3). python’s pyre module and prob2b are used to perform a form and mc analysis. when the failure mechanisms are tuned to the target β-value, the partial factors can be calculated (section 4). [4] 2.2. typical sft-structure the design made for the bjørnafjorden in norway concerns the typical failure mechanisms of an sft structure, with a certain depth and size, and is therefore chosen as reference in this paper. the structure consists of two identical concrete tubes with steel reinforcement. between the two tubes, there are crossbeams and diagonals. the diameter of the tube is not the same along its length, since there are emergency lanes every 250 m. the tube elements are 200 m, and the total length of the sft is approximately 5400 m. the tube is stabilized by steel tethers attached to the seabed with drilled and grouted rock anchors. for this structure, typical loading types were defined, e.g. the permanent downward loads, the buoyancy force, the traffic loading and the wave-current loading (figure 2). for axial loading, the hydrostatic load is considered. the vortex-induced vibrations, temperature loading and creep and shrinkage turned out to be negligible. [4] 2.3. decomposition of the system the entire structure of the sft can be seen as one system. we need to make a decomposition of this system in order to assess the reliability on elementlevel, conform eurocode en1990 [2]. figure 3 shows a decomposition in multiple sft-sections, where each section can fail separately. one sft-element is considered here as one single tube. the framework in between the tubes is not 199 franka e. m. swaalf, raphaël d. j. m. steenbergen acta polytechnica ctu proceedings figure 3. fault tree for an sft figure 4. fault tree for an sft-element included in the calculations, since we only look at the most important force actions. the tube is attached to the bottom of the sea with tethers [1]. we assume four tethers per mooring for our model input. on element level, multiple relevant failure mechanisms can be defined, as can be seen in figure 4. from design experience, the following four mechanisms can be assumed as most important: yielding (1), slackening (2), failure of the tube in longitudinal direction (3) and shear failure of the tube in transverse direction (4). more about these mechanisms will be explained in section 3. compression in transverse direction was first assumed to be an important failure mechanism as well. however, it is proven that the compressive stress is small compared to the concrete resistance. within the scope and planning of this study, it has been assumed that the additional failure mechanisms of corrosion, fatigue, geotechnical failure and accidental failure are not governing due to their complexity and expected research time needed. for geotechnical failure extensive research into the geolocation should be performed [4]. 3. reliability of main failure mechanisms in this section, the four most important failure mechanisms are explained and their limit state functions are formulated. the input parameters are defined by their mean values and standard deviations, and can all be found in table 1. a β-value of 3.8 for a reference period of 50 years is chosen for the individual failure mechanisms, mainly since general partial factors from eurocode en1990 are also determined on a β-value of 3.8 with a reference period of 50 years [2]. looking at the fault trees in figure 4 and figure 3, we expect the system β-value to be lower than the element β-value. however, this subject is not further studied in this paper [4]. 200 vol. 36/2022 submerged floating tunnel design for safety 3.1. yielding of the tethers when the stress in the cross-section of the tether reaches the yield point, the steel begins to deform plastically. the yield point is the point where nonlinear (elastic and plastic) deformation begins. yielding of the tether can occur when the buoyancy force increases or the downward force decreases. this change of loading can take place due to erosion, temporary replacements, shortage of ballasting, or due to waves. the yield strength of steel and the cross-sectional area determine the strength of the tether. for this study, not all tethers need to be investigated separately because equal properties are assumed. both the models for resisting forces and load contain uncertainty, because the used mathematical models are not a full representation of reality. this is included in model factors for uncertainty (θr and θs ). z = θr·fresistance,steel−θs ·(fbuoyancy +fwave−current − fconcrete − fballast − fasphalt− fequipment − fmarine) (1) the resistanceand strength parameters can be plotted in a joint probability density function (5, left). the failure point is the point where the line z = 0 and the contour plot of the function intersect (5, right). from this contour plot, the design points and α-values can be derived. for all four mechanisms, these plots have been assessed [4]. 3.2. slackening of the tethers slack means that there is no tension in the cable anymore, and the stiffness is zero. as a result of slackening, snap forces occur, which can lead to structural failure of the system. due to marine accumulation, extra ballasting, or leakage, slackening can occur. loosening of the anchorages can also occur due to waves, and often lasts for one to two seconds during each wave run. in order to prevent slack, an equilibrium should be maintained between upward and downward forces. in the governing situation for slackening, lift force acts as a downward force. the accompanying drag force has no vertical component, so this force is not included in the formulation [4]. z = θr·fbuoyancy +θs ·(−fwave−current−fconcrete− fballast − fasphalt − fequipment − ftraf f ic − fmarine) (2) 3.3. longitudinal failure of the tube in the longitudinal direction, the tube can deform due to tidal differences or excitation due to waves or currents. furthermore, sea level rise and extreme water levels can cause larger hydrostatic loads on the cross-section, which increases cross-sectional forces and moments. the point along the cross-section with the largest bending moment needs to be further investigated. for this point, two cases need to be investigated: • lower limit: the elastic moment with zero tension in cross-section • upper limit: the plastic moment capacity for an elastic stress state, all stress distributions are considered to be linear. the stress at the bottom of the cross-section can be calculated according to equation 3. σc,total = − next ac − np ac − np · e(x) w + mext(x) w (3) w = π · (d4 − d4) 32 · d (4) where: x = longitudinal coordinate [m] σc,total = total stress in the cross section [kn/m2] mext = applied bending moment [kn m] next = normal force due to hydrostatic pressure [kn ] w = section modulus [m3] np = normal force due to prestressing [kn ] ac = concrete area of the tube [m2] e = assumed eccentricity of the post-tensioning [m] d = outer diameter [m] d = inner diameter [m] it is assumed that the resulting compressive stress in the concrete cross-section is within the acceptable limit. the cross-section should not be subjected to tensile stresses. the limit state function for this mechanism is stated in equation 5. z = θσr · ( np ac + np · e w ) − θσs · ( next ac + mext w ) (5) in reality, a larger external moment can be applied to the cross-section before it fails in terms of cracking and transmitting the load. the cross-section can be seen in figure 6. in any case, the structure fails when the moment exceeds the plastic moment capacity. for the maximum plastic moment capacity, it can be assumed that a third of the cross-section is yielding with an inner arm of 0.8·d [6]. mplastic,max = 1 3 · ac · fy · 0.8 · d (6) the plastic moment at this point is: mplastic = 1 16 · qtotal · l2 (7) the following limit state results in the upper limit for longitudinal failure: [4] z = mplastic,max − mplastic (8) 2norm= normal distribution, log= lognormal distribution, gum= gumbel distribution, det= deterministic value 201 franka e. m. swaalf, raphaël d. j. m. steenbergen acta polytechnica ctu proceedings figure 5. joint probability density function (left) and contour plot (right) for yielding parameter distr. 2 mean st.dev. cv source cross-section tether (at) [m2] norm 0.67 0.017 0.025 [1] yield strength s235 (fy ) [n/mm2] log 285 20 0.07 [5] diameter of the tube (d) [m] norm 15.0 0.375 0.025 [1] thickness of the tube (t) [m] norm 0.85 0.021 0.025 [1] length of the tube (ltube) [m] det 150 [1] gravitational acceleration (g) [m/s2] det 9.81 unit weight of concrete (γc) [kn/m3] norm 24.5 1.7 0.07 [6] unit weight of water (γw ) [kn/m3] norm 10.035 0.4 0.04 [1] drag coefficient (cd ) [-] log 0.7 0.2 0.3 [7] lift coefficient (cl) [-] log 0.1 0.02 0.2 [7] wave-current velocity for 50 years (uc) [m/s] gum 1.5 0.15 0.1 [1] structural weight of asphalt (qasphalt) [kn/m] norm 28 2.8 0.1 [1] weight of permanent equipment (qequipment) [kn/m] norm 10 1 0.1 [1] weight of average solid ballast (qballast) [kn/m] norm 100 10 0.1 [1] traffic load for 50 years (qtraf f ic) [kn/m] gum 50 7.5 0.15 [8] marine load (qmarine) [kn/m] norm 10 2 0.2 [1] area of strands (astrand) [mm2] det 150 [9] initial strength of prestressing y1860 (fs) [n/mm2] log 1300 90 0.07 [5] concrete compressive strength c45/55 (fc) [n/mm2] log 53 8 0.15 [6] model uncertainty for shear capacity θ1 [-] log 1 0.2 0.2 [6] model uncertainty for resistance θr [-] log 1 0.05 0.05 [7] model uncertainty for load effects θs [-] log 1 0.05 0.05 [7] model uncertainty for the capacity of the cross-section θσr [-] log 1 0.05 0.05 [7] model uncertainty for stresses in the cross-section θσs [-] log 1 0.05 0.05 [7] table 1. input parameters for reliability analysis parameter yielding slackening longitudinal failure shear failure β 3.8 3.8 3.8-6.7 5 table 2. the β-values per failure mechanism 3.4. shear failure of the tube a large shear force can occur at the attachment points of the tethers to the tube, due to wave-current forcing. this can lead to shear failure. it turned out that the limit state for shear failure resulted in a β-value of 5. the β-value of shear failure could be tuned near 3.8 by changing the intermediate distance of the tethers. however, it is better to have longitudinal failure first, since this is a ductile mechanism. since the probability of failure due to shear is now significantly lower 202 vol. 36/2022 submerged floating tunnel design for safety figure 6. lay-out of reinforcement: normal reinforcement in green and eccentric reinforcement in red than failure due to one of the other mechanisms, this mechanism is not taken into account for the partial factor analysis [4]. 4. results and discussion 4.1. interpretation in terms of β’s and α’s the resulting β-values from the four failure mechanisms can be found in table 2. the β-value for yielding is a slight underestimation of the real β-value. this is caused by two factors: the traffic load was not taken into account (1) and the wave-current force was only taken as unfavorable loading (2). the β-value for slackening resulted in 3.8, which is also a slight underestimation, since the lift force was only taken as unfavorable loading. the β-value for longitudinal failure turned out to have a value between 3.8 (lower limit) and 6.7 (upper limit). the sensitivity factors (α) are presented in table 3. only the parameters with an α-value larger than 0.1 are shown. the diameter and thickness have two α-values, due to the uncorrelated and correlated calculation respectively. the diameter is dominant over the thickness, and the parameters serve as loading parameters for yielding and as resistance parameters for slacking. the α-value of d decreases significantly for the correlated case. the largest α-values are found for the concrete density, water density, yield strength of the steel tethers, the diameter of the tube and the model uncertainties. the variable loads, i.e. traffic load and wind-current load, have relatively low α-values. most α-values are smaller than either 0.8 (dominant resistance parameter) and −0.7 (dominant loading parameter) or 0.28 (other resistance parameters) and −0.32 (other loading parameters) from eurocode en1990 [2]. this means that the α-value multiplied with β-value results in a smaller distance between the design point and the mean value than eurocode en1990, which consequently leads to a smaller partial factor. in a form analysis, no distinction is made upfront between dominant and non-dominant parameters. furthermore, when more parameters are added, the lower the α-values become, because the sum of the α-values squared always adds up to one [4]. 4.2. interpretation in terms of γ’s for the assessment of partial factors, the parameters with α-values above 0.1 are chosen stochasts, and the rest of the values as deterministic parameters. consequently, the design points of the stochasts are calculated. parameters are divided into favorable load, unfavorable load and variable load. then, characteristic values according to eurocode en1990 were determined [2]. the partial factors of the full probabilistic design can be calculated by using this design value and characteristic value. eurocode en1990 distinguishes in types of calculations [2]. for a strength calculation, the strconditions can be used. these factors can thus be compared with the calculated factors from tether yielding and longitudinal tube failure. for an equilibrium calculation, the equ-conditions can be used. these factors can be used for slackening of the tethers. the factors according to the strand equ-conditions and the calculated partial factors, according to the probabilistic design, can be found in table 4. for the str-criteria, the favorable permanent load factor and the material resistance factor fit well. however, the general partial factors for the unfavorable permanent load and for the variable load seem to be very conservative. based on this case study, the general partial factors for variable loading can be decreased with at least 20 %. since the influence of variable loads on the structure is small, the decrease 203 franka e. m. swaalf, raphaël d. j. m. steenbergen acta polytechnica ctu proceedings parameter α yielding α slackening α longitudinal failure fy [n/mm2] 0.50 γw [kn/m3] -0.48 0.43 -0.52 γc [kn/m3] 0.47 -0.52 0.52 fs [n/mm2] 0.55 d [m] -0.46, -0.04 0.27, 0.11 -0.35, -0.11 t [m] 0.13, -0.34 -0.19, 0.08 0.14, -0.21 θr [-] 0.32 0.48 θs [-] -0.29 -0.48 θσr [-] 0.32 θσs [-] -0.32 table 3. the α-values from form analysis parameter str equ yielding slackening longitudinal failure resistance 1.15 1.07 1.08 unfavorable permanent load 1.35 1.1 1.23 1.19 1 favorable permanent load 1 0.9 0.98 0.82 0.86 unfavorable variable load 1.5 1.05 1.18 1 1.04 table 4. summary of the resulting partial factors from the models of this factor will not result in significant changes in the design. thus, the economic advantage is small. in contrast, a change in water density will have a large influence on the design. thus, decreasing the general partial factor for unfavorable permanent load will have a significant effect on the design. according to this case study, the factor of 1.35 could be decreased with almost 10 %. however, conclusions cannot be too firm, since this calculation was performed for one case only. for the equ-criteria, partial factors from eurocode en1990 are not safe to be applied. the partial factor for unfavorable permanent loading is insufficient. this factor should be increased by 10% in order to design a sft [4]. 5. conclusions this paper showed that the reliability requirements of the sft can be met in the design, and that the design can be optimized by a full probabilistic calibration of partial factors. the following conclusions were drawn: • the conventional reliability methods (i.e. form and mc) can be adopted in a reliability based design of an sft. permanent loads proved to be dominant, i.e. concrete and water density, and thus significantly influence the design of the sft. the same reliability methods can be applied to other sfts, but they might result in different designs based on geolocation specific circumstances. • a target reliability is required to perform the reliability analyses. the general partial factors from eurocode en1990 are based on a β-value of 3.8 for a reference period of 50 years [2]. therefore, in this research, a target β-value of 3.8 for 50 years was chosen as starting point for the individual failure mechanisms. design parameters were determined so that this target value was met. the actual βvalues were computed using simplifications that are by definition safe. on system level, it is expected that the β-value is lower than 3.8, but this is not studied further. • four important failure mechanisms for the sft were derived using a decomposition of the system. these important mechanisms are yielding of the tethers, slackening of the tethers, longitudinal failure and transverse shear failure of the tube. it has been assumed within the scope and planning of this study that the additional failure mechanisms of corrosion, fatigue, geotechnical failure and accidental failure are not governing due to their complexity and expected research time needed. for geotechnical failure extensive research into the geolocation should be performed. slackening proved to be the governing failure mechanism over the other three mechanisms. the resistance of slackening depends on the force equilibrium, whereas the resistance of the other mechanisms depends on structural strength. • the influence factors (α-values) from the form analysis indicated that the most dominant parameters were the concrete density, water density, yield strength of the steel tethers, the diameter of the tube and the model uncertainties. the resulting α-values from form do not only depend on the coefficient of variation, but also on the absolute value of the mean of the parameter. the permanent loading parameters turned out to be dominant, because their relative contribution to the total load is large. the variable loads, i.e. traffic load and wind-current load, have relatively low α-values. • the general partial factors from eurocode en1990 204 vol. 36/2022 submerged floating tunnel design for safety were never calibrated on sft-type of structures [2]. they seem nevertheless safe enough to apply for the strength (str) cases (table 4). the general partial factors for the unfavorable permanent load and for the variable load turned out to be very conservative. a decrease of 20% for the variable loading factor and 10% for the factor of unfavorable permanent loading is proposed. since the influence of variable loads on the structure is small, the decrease of this factor will not result in significant changes in the design. in contrast, a change in one of the permanent loading parameters will have a large influence on the design. thus, decreasing the factor for unfavorable permanent load will have a significant effect on the design and is economically attractive. • the equilibrium (equ) criteria from eurocode en1990 should be applicable to the slackening mechanism. however, the general partial factor for the unfavorable permanent loading turned out to be insufficient (table 4). this factor should be increased by 10% in order to be able to design a safe enough sft [4]. references [1] reinertsen, o. dr. techn. olsen, norconsult. bjørnafjord submerged floating tube bridge 2016. [2] en1990. eurocode 0: basis of structural design 2011. [3] m. baravalle, j. köhler. risk and reliability based calibration of design codes for submerged floating tunnels. procedia engineering 166:247–254, 2016. https://doi.org/10.1016/j.proeng.2016.11.547. [4] l. swaalf. reliability analysis of submerged floating tunnels 2020. [5] j. melcher, z. kala, m. holický, et al. design characteristics of structural steels based on statistical analysis of metallurgical products. journal of constructional steel research 60(3-5):795–808, 2004. https://doi.org/10.1016/s0143-974x(03)00144-5. [6] en1992. eurocode 2 part 1-1: design of concrete structures: general rules and rules for buildings 2005. [7] jcss. probabilistic model code. joint committee on structural safety 2001. [8] en1991-2. eurocode 1 part 2: traffic loads on bridges 2015. [9] tensabv. multi-strand post-tensioning system 2007. 205 https://doi.org/10.1016/j.proeng.2016.11.547 https://doi.org/10.1016/s0143-974x(03)00144-5 acta polytechnica ctu proceedings 36:198–205, 2022 1 introduction 2 methodology 2.1 procedure 2.2 typical sft-structure 2.3 decomposition of the system 3 reliability of main failure mechanisms 3.1 yielding of the tethers 3.2 slackening of the tethers 3.3 longitudinal failure of the tube 3.4 shear failure of the tube 4 results and discussion 4.1 interpretation in terms of 's and 's 4.2 interpretation in terms of 's 5 conclusions references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0015 acta polytechnica ctu proceedings 34:15–19, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague a modular extension of a finite element contact implementation ondřej faltus∗, martin horák czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: ondrej.faltus@cvut.cz abstract. the structural mechanics module of the oofem finite element software has been developed to include various algorithms for solving contact mechanics problems. after developing small strain contact algorithms in the 2d domain and their extension to large strain applications [1, 2], the present contribution extends the existing framework to the 3d domain. after reviewing the current code and comparing existing solutions found in the literature, we identify the common ground between 2d and 3d applications, then propose and implement the necessary changes and additions to smoothly integrate the 3d support into existing code. tests and example problems are presented to confirm the functionality of the resulting implementation. keywords: contact, large strains, penalty method, 3d contact, node-to-segment, oofem. 1. introduction contact mechanics, as a specific subdiscipline of structural mechanics, lend themselves very handily to computational implementations. analytical solutions to contact mechanics problems are scarcely found. since the end of the 19th century, when heinrich hertz studied the contact of convex elastic bodies without friction [3], progress in this originally fringe discipline had been very slow until the advent of computational methods, mainly the finite element method (fem) in the latter half of the 20th century [4, 5]. today, more and more complicated problems can be studied with the help of more and more involved algorithms, owing to the technological advances in computational hardware. in our past work [1, 2], we have developed the basis of contact algorithms in the oofem software. oofem [6] is a finite element computational software package developed at the czech technical university in prague. from a programming perspective, oofem is built as a strictly object-oriented c++ codebase. in our past implementations, we have paid close attention to this fact, adapting some of the commonly in literature presented programmatic solutions to suit the object-oriented philosophy better. so far, contact algorithms in oofem can handle contact problems in the 2d domain, using the node-to-node and node-tosegment discretizations and alternatively the penalty method or lagrangian multiplier method of contact condition enforcement. of these options, the penalty method node-to-segment approach is adapted for use with geometrically nonlinear problems as well. this work aims to extend our previous implementation to the 3d domain. keeping in with the idea of object-oriented code, we try to achieve this by extending the existing objects and algorithms they represent instead of coding everything anew. chiefly, we would like to avoid the "contact element" approach often presented in literature [4, 7, 8] and create instead a universal contact condition object able to interact with every possible contact segment type. 2. theoretical basis at the core of our modular code extension lies the desire to adapt the existing geometrically nonlinear penalty contact formulation for 2d domains to fully work in the 3d domain as well. for this purpose, we would like to derive a universal formulation for the forces and stiffness applicable to all cases equally. 2.1. formulation of penalty contact the penalty formulation counts as one of the most straightforward possible implementations of contact conditions into the fem framework. its governing idea is the introduction of an arbitrary penalty parameter to the force vector, aimed at penalizing unwanted states of penetration among the contacting bodies. in other words, we express the force on the contact boundary as h(−gc)pgc, where p is the chosen penalty parameter and gc the contact gap (negative in penetration) [1, 4, 5, 7]. h is the heaviside function. to this end, the energy contribution of the contact process can be expressed as w c(u) = ∫ γc 1 2 h(−gc(u))pg2c (u) dγ (1) where u is the global vector of displacements and γc is the contact boundary. note that for a node-tosegment contact discretization, the integral morphs into a sum over a finite set of discrete contact points. with the first variation of the contact energy, we obtain the internal force vector as 15 https://doi.org/10.14311/app.2022.34.0015 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en ondřej faltus, martin horák acta polytechnica ctu proceedings fc = ∫ γc pgc ( n ∗t n ||n|| ) dγ (2) where n is the normal vector to the contact boundary in the deformed configuration and n ∗ is the extended n-matrix: a matrix serving to distribute the forces to the according degrees of freedom of both the external node and the nodes of the element forming the contact surface (see [1] for a more detailed discussion of the concept). under the assumption of small strains, the normal vector and the extended n-matrix would remain independent of u, giving with the second variation of contact energy a simple formula for the algorithmic stiffness contribution [1, 2] kc = ∫ γc p ( n ∗t n ||n|| ) ( n ∗t n ||n|| )t dγ (3) however, those assumptions no longer hold with nonlinear geometry, and a more complex stiffness formula has to be derived. keeping in mind the desired universality of the solution for both 3d and 2d domains, we shall do that in the following paragraph. 2.2. a universal stiffness formula in expression (2) for the internal forces of contact, the gap function gc depends on the current deformation state of the domain, with the variation δgc = n tv δd (4) with nv = n ∗t n/||n|| and δd being the variational displacements of the related degrees of freedom. moreover, the extended n-matrix n ∗ and the unit normal vector ν = n/||n|| depend on the convective natural coordinates ξ denoting the contact point on the element surface. the β-th natural coordinate ξβ has the variation δξβ = a−1αβ (tv,α − gcbv,α) δd = dv,β δd (5) where: bv,α = b∗tα ν tv,α = n ∗t tα (6) the matrix b∗α being the derivative of n ∗ with respect to ξα, the vector tα being the tangent vector to the surface in the α-th direction and a being an extended metric tensor of the surface, composed of the covariant metric tensor m and the curvature tensor κ [4] aαβ = mαβ − gn καβ (7) summation convention over α and β (parametric surface coordinate indicies) is implied wherever applicable. after substitution of these variations into the known expressions and exploitation of the orthogonality of the normal and tangent vectors, we can express the desired tangential stiffness as kc = ∫ γc pn tv nv (8) + pgc ( bv,αd t v,α + dv,αb t v,α + καβ dv,β dtv,α + gcm αβ b̄v,αb̄ t v,β ) dγ where b̄v,α = bv,α + καβ dv,β (9) and mαβ is the contravariant metric tensor (i.e. m−1). note that the first term in equation (8) corresponds perfectly to the geometrically linear formula (3). this formula has been derived with the express aim to make no assumptions about the dimension of the domain space or the number of nodes on the element boundary. it is, therefore, a universal formula for all cases of nonfrictional contact. letting α and β only assume the value of 1, setting m to element length squared and κ to 0, an expression for a linear element in 2d space is obtained, which reduces perfectly to the stiffness formula found for this case in [4] and used in [2]. 3. implementation this section shall discuss the precise intricacies of the chosen approach to implementing the contact algorithms. 3.1. the contact condition and contact segment objects contact usually occurs among two bodies within the domain space, of which at least one is deformable and described in fem by a mesh of finite elements. by definition, when the contact gap is open, no solid matter occupies the area between the two bodies, and therefore no mesh is present. from a practical programmatic perspective, an essential part of any fem algorithm is the assembly process. in this process, all the elements within the domain space contribute to the global stiffness matrix and the global vector of internal forces, assembling their contributions to their proper places. in the case of contact, it is unclear how to introduce contact forces and resulting algorithmic stiffness into this process. in literature [4, 5, 7], it is customary to introduce so-called contact elements. those are finite elements occupying the physically empty space in between the contacting objects. those elements perform the aforementioned force and stiffness assembly function. therefore, the gap area in the reference configuration has to be meshed with those special elements describing the contact algorithm. this approach has several noticeable drawbacks. firstly, in the node-to-segment approach, for example, a separate contact element has to be formulated for 16 vol. 34/2022 a modular extension of a finite element contact implementation figure 1. a uml diagram describing the concept of object relationships within oofem. each possible case of element edge or surface forming the segment (in the segment-to-segment approach, matters would be even much worse). for a codebase with a reasonable claim to universality, this could shortly result in a significant increase of necessary code. secondly, the formulation of those elements may be restrictive in terms of the possible contact cases. a contact element usually only pairs one node with one boundary segment. each node deemed possible to come in contact with a given segment, therefore, requires a separate element defined, and this has to be decided already at the time of meshing. already in the existing implementation in oofem, we have diverged from this standard practice. in line with object-oriented programming principles, the aforementioned assembly process in oofem is handled with the use of objects. apart from object classes representing elements, another class capable of contributing to the global stiffness matrix and force vectors is the activeboundarycondition class. we have chosen to handle contact algorithms through subclasses descended from that. the principle is illustrated diagrammatically in figure 1. all oofem objects are collectively stored in the overarching domain class. a node2segmentpenaltycontact object is a subtype of an activeboundarycondition, which itself is a subtype of a boundarycondition. this contact condition class maintains a generalized reference to a contactsegment object, which, in the case it is of the boundarycontactsegment subtype, can reference element objects relevant for its task of describing an element boundary. this way of structuring the code offers us a great potential of generality in large parts of the code and presents a more intuitive way of modeling the real physical idea of the contact phenomenon. 3.2. universality of the contact condition object in our previous work, four boundary condition objects have been created, namely node2nodepenaltycontact, node2nodelagrangianmultipliercontact, node2segmentpenaltycontact, and node2segmentlagrangianmultipliercontact [1, 2]. in the case of the node-to-node approach, the contact condition object itself handles the references to the nodes concerned, requiring no interaction with any other objects. its universality and extension to the 3d domain is therefore trivial and not a subject of this paper. in the node-to-segment case, however, the operations within the code rely heavily on an interaction between the condition and an object representing the contact segment. as described by figure 1, we aimed to maintain the connection of the boundary condition class only to the generalized contactsegment parent class, rather than differentiating among its subtypes, which include formulations as diverse as an analytical polynomial function in 2d space and a segment representing multiple 3d element surfaces. the formulas necessary to this universal approach have already been discussed in section 2.2. the only necessary part of the implementation part is to ensure the ability of the code to handle all possible matrix sizes of the various contributions. extensive use has been made of the std::list array object in the c++ language, which allows for an unspecified number of objects to be passed between c++ functions. to this end, the functions in the 2d contact segment have been altered to return an array of size one (of, i.e., tangent vectors) as opposed to a single object. on the level of the node2segmentpenaltycontact class, the code is written to always loop through those arrays regardless of their size. key to the computations in the node2segmentpenaltycontact class are the computetangentfromcontact() and computeexternalforcesfromcontact() methods, which compute the stiffness and force contributions, respectively, for a given node-segment pair. in addition to those, auxiliary methods exist to compute the various expressions seen in formula 8, e.g. the computenvmatrixat() method for the nv matrix or the computemodifiedbvmatricesat() method for the (one or multiple) b̄v matrices. those then poll the contact segment for the relevant information, such as gap value, metric and curvature tensor, normal and tangent vector(s), the n matrix etc. 3.3. the universality of contact search algorithm another issue with code universality arises specifically in the 3d domain. contrary to the 2d case, in 3d, even the limitation to finite elements with linear formulation does not necessarily ensure the linearity of their boundary [9]. more specifically, only surfaces of triangular shape can be described linearly. we have come to the conclusion that limiting our implementation to those would severely hamper the usability of the resulting code; it was, therefore, necessary to take the possible nonlinearity of the contact surface into account when conducting a contact point search. the solution 17 ondřej faltus, martin horák acta polytechnica ctu proceedings (a) (b) (c) figure 2. node-to-segment contact with penalty formulation in the 3d domain: tests on a linear wedge and a linear brick element. displayed geometries correspond to the 0th (undeformed configuration), 40th, and 60th steps of the computation. coloring represents the vertical normal stress σyy . a) a test on a linear wedge element. b) a test on a linear brick element, with the initial contact point at an axis of symmetry of the top surface. the deformed state remains symmetric along this axis. c) a test on a linear brick element with an arbitrary initial contact point. the apparent gap is only a consequence of the imprecise rendering of the resulting bilinear surface. is an implementation of a closest-point-projection algorithm in the computecontactpoint() method of the linear3delementsurfacecontactsegment class. the parametric coordinates of the desired contact point (i.e., the point on the surface closest to the contacting node, regardless of whether contact occurs) can be expressed as a vector ξ minimizing the distance function f (ξ) = 1 2 (r − ρ(ξ)) · (r − ρ(ξ)) (10) where r are the global coordinates of an external point (the node), and ρ are the global coordinates of the contact point (on the surface). solving this minimization problem with the newtonraphson scheme leads to a convenient expression for the iterative increment δξ in the form δξ = −a−1 df dξ (11) where a is the already well-established extended metric tensor of the contact surface (see (7)). 4. computational examples in light of the original aim of this paper, the extension of oofem’s contact formulations to the 3d domain, we shall demonstrate the functionality of the described implementation on several 3d test examples. the tests concern the node-to-segment formulations for geometrically nonlinear cases. in one case, the contact segment is made of a triangular surface of a linearly formulated wedge element, while in the other two cases, the used element is a linear brick. therefore, the first element represents an example of a fully linear surface without curvature, while the latter concerns a bi-linear surface, where the curvature tensor is generally nonzero. on the brick element, two different tests are performed; in the first one, the initial contact point is positioned on one of the axes of symmetry of the surface, in the other one, an arbitrary point is chosen. the elements in question are the oofem elements lwedge and lspace, respectively. above them, a single ltrspace element is positioned. this is a linear tetrahedral element, which serves, however, only to provide the contacting node. all its nodes are fixed, and the lower node is loaded by a prescribed vertical displacement. this node is designed to come into contact with the top surface of the lower element and push it downwards. all nodes of the contact surface are free; the lower element is fixed by dirichlet boundary conditions on its lower base. all elements in these test simulations have a linear 18 vol. 34/2022 a modular extension of a finite element contact implementation elastic material model assigned, with young’s modulus and poisson’s ratio e = 500 mpa and ν = 0. in oofem, such a material model translates to the saint venant-kirchhoff law in the case the geometrical nonlinearity of the element is turned on [6]. loading is performed in 60 steps in each case. the total displacement prescribed for the external node is 0.3 m, i.e. 0.005 m per loading step. the results of the test are pictured in figure 2. for the visualization of the results, paraview, an open-source software package, has been used [10]. we can observe that the deformation of the surface follows the positioning of the contact point. for the brick element, in the first case, the contact point remains on the axis of symmetry, and the rotation of the contact surface only follows the other axis, while in the second case, the entire surface twists and assumes a bi-linear deformed shape. note that the somewhat apparent gap between the element surface and the contacting node is only a result of the rendering software’s inability to portray this bi-linear surface properly. 5. conclusions the paper describes an object-oriented approach to handling contact mechanics in fem. as a result of this object-oriented approach, a universal contact algorithm is necessary. this is derived on a theoretical basis, and the implementation in an open-source finite element code oofem is then discussed in detail. to verify the practical usability of the chosen approach, tests are performed on contact in the 3d domain, which is a feature previously unavailable in the existing oofem contact implementation. acknowledgements the authors would like to acknowledge the support received for work on this project from a ctu grant no. sgs21/037/ohk1/1t/11. references [1] o. faltus. object-oriented design and implementation of contact mechanics into finite element code oofem. master’s thesis, czech technical university in prague, 2020. [2] o. faltus, m. horák. finite element implementation of geometrically nonlinear contact. acta polytechnica ctu proceedings 30:18–23, 2021. [3] h. hertz. on the contact of elastic solids. z reine angew mathematik 92:156–171, 1881. [4] p. wriggers. computational contact mechanics. springer, new york, 2nd edn., c2006. [5] v. a. yastrebov. numerical methods in contact mechanics. wiley, hoboken, nj, 2013. [6] b. patzák. oofem home page, 2000. http://www.oofem.org. [7] a. konyukhov, r. izi. introduction to computational contact mechanics. wiley, chichester, west sussex, 2015. [8] t. a. laursen. computational contact and impact mechanics: fundamentals of modeling interfacial phenomena in nonlinear finite element analysis. springer science & business media, 2003. [9] o. c. zienkiewicz, r. l. taylor. the finite element method. butterworth-heinemann, boston, 5th edn., 2000. [10] u. ayachit. the paraview guide: a parallel visualization application. kitware, 1st edn., 2015. 19 acta polytechnica ctu proceedings 34:15–19, 2022 1 introduction 2 theoretical basis 2.1 formulation of penalty contact 2.2 a universal stiffness formula 3 implementation 3.1 the contact condition and contact segment objects 3.2 universality of the contact condition object 3.3 the universality of contact search algorithm 4 computational examples 5 conclusions acknowledgements references my title https://doi.org/10.14311/app.2022.36.0025 acta polytechnica ctu proceedings 36:25–32, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague estimating running safety factor of ballastless railway bridges using tail modelling reza allahvirdizadeh∗, andreas andersson, raid karoumi kth royal institute of technology, division of structural engineering & bridges, department of civil and architectural engineering, brinellvägen 23, 100 44 stockholm, sweden ∗ corresponding author: rezaal@kth.se abstract. excessive vertical acceleration of ballastless railway bridges subjected to vibrations induced by passing trains is one of the governing design criteria for bridges in high-speed lines. however, to the authors’ knowledge, the corresponding design limit is not based on a solid theoretical or experimental background. moreover, the traditionally applied safety factor also suffers from these concerns. therefore, in the present study, a crude probabilistic approach is adopted to evaluate the consistency and reliability of this safety factor. for this purpose, deterministically designed bridges (using conventional methods) with short to medium spans are considered. then, their reliability is evaluated using simulation-based techniques and extreme value theory, i.e., tail approximation. then, the existing safety factor is calculated to evaluate the consistency of the current approaches, and possible new values are proposed based on the desired target reliabilities. keywords: ballastless bridges, high-speed railway bridges, running safety, safety factor, structural reliability, tail modelling. 1. introduction the increase in operating speed of trains in recent decades leads to excessive vibration of infrastructure, especially bridges, which raises new safety concerns. such vibrations can affect the structural performance of bridges, jeopardize the safe passage of trains on bridges (denoted here as running safety), and disrupt passenger comfort. ensuring running safety requires preventing loss of contact between wheels and rail, the occurrence of which increases the risk of train derailment. this phenomenon can be formulated as an unloading ratio, which reads as eq.(1) [1]. ∆p pst = pst − pmin,dyn pst (1) where pst is the static vertical wheel load and pmin,dyn is the minimum dynamic vertical wheel load. theoretically, contact loss occurs at a unloading ratio of 1, but [1] limits this to 0.6. considering the formulation presented, the study of the running safety problem must take into account various subsystems, namely the passing train, the rail, the track, the bridge components, the boundary conditions, and the interaction between all these subsystems. the train and the track are coupled to each other by the interactive wheel-rail force, which is transmitted to the bridge through the track. it is worth noting that there are a variety of track types in the world, which can generally be categorized into two classes: ballasted and ballastless (also referred to as non-ballasted or slab track). for the first type, a granular material is formed between the sleepers and the bridge deck. for the latter, however, there are different construction methods, about which the interested reader can find a detailed overview in [2], but in general they consist of two concrete layers and a bituminous mixture in between. clearly, the use of such complicated computational models requires a high level of expertise and involves expensive computational costs. therefore, for practical purposes, moving load models are often used, which are computationally very efficient. in addition, it has been shown that such reduced models can lead to acceptably accurate answers when the train weight is much smaller than the bridge (which seems to be the case for passenger trains) and also in cases where the study of the dynamic behavior of trains is not sought [3]. this approach cannot model the contact loss between wheels and rail. therefore, the design rules implicitly control the ride safety criterion by instead limiting the vertical acceleration of the bridge deck [4] or its deflection [5]. this assumes that contact loss occurs at accelerations greater than 1.0 g; however, the authors found no theoretical or experimental justification for this assumption. the relationship between this criterion and the unloading ratio (wheel-rail contact loss) was previously studied using train-track-bridge interaction (ttbi) models, where it was found that the acceleration criterion almost always dominates the design [6, 7]. in addition, an arbitrary safety factor of 2 is applied to the above threshold. therefore, the maximum allowable vertical acceleration of ballastless railroad bridges should be limited to 5 m/s2. this safety factor is borrowed from the phenomenon of ballast instability [4]; however, it should be emphasized that the consis25 https://doi.org/10.14311/app.2022.36.0025 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en r. allahvirdizadeh, a. andersson, r. karoumi acta polytechnica ctu proceedings tency and reliability of this safety factor is questioned even for bridges with ballast [8]. considering this situation, a rough probabilistic approach is taken in this article to evaluate the reliability of the conventional safety factor for ballastless railroad bridges. the rest of this article is divided into the review of the applied reliability assessment method in 2, the description of the constructed computational models in 3, the formulation of the reliability problem in 4, and the estimation of the corresponding safety factors in 5. the article concludes with a summary of the main results. 2. reliability assessment using tail modelling the probability of failure (pf ) is generally estimated by calculating the presented multiple integral in eq. (2): pf = p(g(x) ≤ 0) = ∫ d fx(x)dx (2) where x ∈ rm is the vector of basic random variables, fx(x) is their joint distribution probability, and g(•) is the limit state function. the latter is basically a decision boundary separating the failure domain (d) from the safe region. these regions are distinguished based on the sign of the limit state function, i.e., failure occurs when g(x) = r − s ≤ 0; where r is the capacity corresponding to the failure mode and s is the demand (e.g., load, displacement, stress, or strain). at this point, it should be mentioned that reformulating the reliability problem as a ratio between action and corresponding capacity, i.e., g(x) = s/r instead of subtracting them, changes the failure domain to the region where g(x) ≥ 1. the calculation of this integral is usually not tractable in real applications. this is because the joint probability distribution of the basic random variables and the failure domain are generally not available. therefore, a variety of methods have been developed to overcome these problems. the best known of these are crude monte-carlo simulation (mc) and first and second order reliability methods (form /sorm). the crude mc gives the most accurate estimate of the failure probability, making it the most widely applicable approach. it can be expressed as eq. (3): pf = ∫ i[g(x) ≤ 0]fx(x)dx = e [ i[g(x) ≤ 0] ] ≈ 1 n n∑ i=1 i[g(x̂i) ≤ 0] (3) where i(•) is an indicator function equal to one if true and zero otherwise. also, n is the number of realizations (samples) of the basic random variables. although it is generally applicable, crude mc is not an efficient approach for the majority of reliability problems. this is because its accuracy (represented by the coefficient of variation of the estimated probability of failure) is inversely proportional to the square root of the number of samples (∝ n−1/2) and also to the order of the probability of failure. considering that the failure probability in structural engineering is usually in the range of 10−4 − 10−3, performing mc incurs significant (and usually unaffordable) computational costs for reliability assessment tasks. moreover, form is basically an optimization-based solution that finds the smallest distance between the point with the largest likelihood on the transformed joint probability distribution in the standard normal space and the most probable point (mpp). the latter should be on the limit state function approximated by the first taylor expansion around the mpp.therefore, form reads as eq. (4). β = arg min t 1 2 tt t s.t. g(t) = 0 (4) where β = −φ−1(pf ) is the safety index, t = t(x) is the transformed basic random variables transformed into the standard normal space, and g(•) is the reformulated limit state function based on the transformed variables. it is worth noting that basic random variables can be transformed into standard normal space, for example, using the nataf transformation [9]. an application of form to evaluate the running safety of ballasted high-speed railway bridges is presented in [8]. as mentioned earlier, this method is computationally efficient but may not be accurate enough for highly nonlinear limitstate functions [10, 11]. considering the discussed issues regarding the efficiency and accuracy of classical reliability assessment techniques, they were later improved by the introduction of more advanced methods. these include important sampling (is) [10, 12], subset simulation (ss) [13], surrogate-based approaches [14, 15], and tail estimation (approximation). the explanation of these methods is beyond the scope of this article; therefore, the interested reader is referred to the references provided. therefore, only the latter approach used in this study is explained here. the failure event is usually a rare event among all possible scenarios of a designed system; therefore, the failure probability is located at the tails of the cumulative distribution function (cdf) of the limitstate function. thus, the basic idea of tail modeling is to approximate the tail of the true distribution by another distribution with an equivalent tail by running a limited number of simulations. the tail equivalence property reads as [16]: lim g(x)→∞ 1 − t [g(x)] 1 − f[g(x)] = 1 (5) where f [g(x)] and t [g(x)] are cdf of the limit state function and that of the tail model, respectively. this equivalence holds for the tails of the true cdf (upper 26 vol. 36/2022 running safety factor of railway bridges using tail modelling or lower ones); therefore, a threshold (denoted here as u) should be chosen to fit the tail model for values larger (or smaller) than this. this results in the fitted tail being formulated based on a transformed variable called the exceedance (z = x − u). the concept of tail approximation, threshold, and exceedance is shown schematically in figure 1. it should be noted here that the limit-state function in figure 1 has been reformulated to show the region of failure at the upper end for illustrative purposes. 0 g(x) f [g (x )] safe failure z = g(x)-u cdf of g(x) fitted tail u figure 1. concept of tail modelling. the most widely used tail model is the generalized pareto distribution (gpd), given as eq.(6). in addition, the sigmoid (eq.(7)) and exponential functions have been used for reliability assessment of high-speed railway bridges in [17], where it was concluded that the sigmoid function can lead to a more accurate estimate of the probability of failure. tz (z) =  1 − 〈 1 + kz/ψ 〉−1/k + k ̸= 0 1 − exp(−z/ψ) k = 0 (6) tz (z) = c1 + c2[ 1 + exp (z/c3) ]c4 (7) where k is the shape parameter of gpd, ψ its scale parameter, ci, i = 1, ..., 4 are constants of the sigmoid function, and ⟨a⟩+ = max (0,a). these parameters can be calculated using either the maximum likelihood method (mle) or the least squares method (regression). in [18], it was shown that using a single-tail model does not guarantee that the failure probability is always estimated accurately. therefore, they proposed to calculate the probability of failure based on a linear regression on the safety index of the top 10% of the data (betalt), a quadratic polynomial fit to the top 50% of the safety index data (betaqh), and a quadratic fit to the logarithmic transformation of the safety index (lnbetaqt) [18]. then they proposed to use the median value of the estimated failure probability from the discussed approaches, showing for benchmark problems that it is almost always at least the second best estimate. therefore, the same approach is followed in this study, using the sigmoid function as the tail model as well. in addition to the assigned parameters of the tail model, its performance also depends on the threshold considered. in [19], it was proposed to consider 1.5 √ n (where n is the total number of data) to fit the tail model (hereafter referred to as hasofer’s method). similarly, [20] suggests using 0.02n and 0.1n for cases of 50 < n < 500 and 500 < n < 1000, respectively (hereafter referred to as boos’ method). furthermore, in [21], they discussed that thresholds close to the central data cause the fitted tail to suffer from high bias, while very large thresholds can lead to tail approximations with high variance ( trade-off between bias and variance). in this context, they proposed to select the optimal threshold by calculating the mean square error (mse) of the estimated safety index (β) at all thresholds considered. the mse can be calculated as the sum of the squared bias and the variance of the estimated safety index at a given threshold (see eq. (8)). given the definition of bias (see eq.(9)) and variance (see eq.(10)), the expected value of the estimated safety index is needed. therefore, they proposed to find the expected value of the safety index using the bootstrap method. an illustrative example of the latter approach is shown in figure 2, where the horizontal axis represents the threshold value as the quantile of all data. it is worth noting that for illustration purposes, the well-known ishigami function is used in the form of eq. (11) to represent the limit state function [22]. threshold selection using this approach is more robust than other methods, although it increases the computational cost. since this additional computational cost is only at the level of post-processing existing data, it would be affordable. mse(β|u) = bias2(β|u) + var(β|u) (8) bias(β|u) = e[β̂(u)] − β (9) var(β|u) = e [ (β̂(u) − e[β̂(u)])2] (10) f(x) = 16.5 − [ sin(x1) + 7 sin2(x2) + 0.1x43 sin(x1) ] xi ∼ u (−π,π) i = 1, 2, 3 (11) 27 r. allahvirdizadeh, a. andersson, r. karoumi acta polytechnica ctu proceedings 0.88 0.9 0.92 0.94 0.96 quantile 0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95 m s e mse best threshold 1 figure 2. threshold selection using mse estimator. 3. computational model as mentioned earlier, this study is concerned with the running safety of ballastless high-speed railway bridges. this problem consists of four subsystems, namely the passing train, the rail, the slab track and the bridge, and their interactions. the used 2d computational model of the simply supported bridges is shown in figure 3. the rail, slab track and bridge are modeled as eulerbernoulli beams. in addition, the rail pads between the rail and the slab track are modeled using lumped springs and dashpots, while the sub grades are accounted for by distributed springs and dashpots. the structure of the train (i.e., the corresponding degrees of freedom) is neglected, resulting in the train being modeled as a series of moving loads. the trains under consideration share a bogie between two adjacent coaches (known as articulated trains) and it is assumed that the axle loads are identical for each realization. since the loads move along the rail, the components of the load vector that do not relate to the rail are zero. therefore, the equation of motion can be expressed as eq. (12) using partitioned matrices. the interested reader can find detailed information about the computational model in [6, 23]. mr 0 00 ms 0 0 0 mb     ÿr ÿs ÿb  +   cr crs 0csr cs csb 0 cbs cb     ẏr ẏs ẏb   +   kr krs 0ksr ks ksb 0 kbs kb     yr ys yb   =   fr 0 0   (12) where m, c, and k represent the mass, damping, and stiffness matrices of each subsystem, respectively. the indices r, s, and b represent rail, slab track, and bridge, respectively. each subsystem is represented by diagonal terms and the off-diagonal partitioned matrices couple the subsystems (rail to slab and slab to bridge). the rail to slab coupling matrices include the lumped spring and the dashpots of the railpads. similarly, the rail to bridge coupling matrices include the distributed spring and the dashpots of the subgrade. in addition, y is the vector of subsystem displacements and fr is the force vector. note that the force vector is time dependent and must be reconstructed at each step based on the position of the train. next, the equilibrium equations of motion are solved using the hilber-hughes-taylor (hht) direct time integration method. then, the calculated responses are filtered using a low-pass filter whose cut-off frequency corresponds to the minimum of the third vibration frequency of the bridge or 30 hz [24]. the constructed computational model cannot capture the effect of unevenness in the rail geometry (known as rail irregularities) on the responses of the bridge. rail irregularities are random in nature and can be modeled by a stationary gaussian process to generate different rail profiles [25]. then they are treated as an additional force term using the wheelrail contact stiffness [23]. in [26], it was shown that rail irregularities can significantly amplify the maximum vertical accelerations of the bridge. therefore, they proposed random amplification factors as an alternative method for modeling rail irregularities. a similar approach is taken by conventional codes, but they consider it as a deterministic variable. this codified amplification factor is considered here; it reads as eq.(13) [24] for the rails with good quality. ria = 1 + α 200 [ 56e−(l/10) 2 + 50( lf1 80 − 1)e−(l/20) 2 ] (13) where α = min (v/22, 1), v is the train speed, l is the length of the bridge span, and f1 is the fundamental frequency of the bridge. it should be noted that the use of ria does not necessarily lead to conservative response predictions [26]; however, this possible drawback is neglected in this study. since the main objective of the article is to evaluate the consistency of the conventional safety factor, the speed of passing trains is limited to the allowable operating range, which is assumed to be 240 km/h here. this is achieved by increasing the operating speed of high-speed trains on the swedish network by 20%, which is about 200 km/h [27]. moreover, only the maximum response of the bridge is needed to estimate the corresponding failure probability. therefore, the maximum vertical acceleration of the deck is calculated only for critical speeds (denoted here as vcr). the critical speed corresponds to the situation where the excitation frequency of the passing train coincides with the frequency of the bridge (resonance phenomenon); this is obtained from eq.(14) [28]. it should be noted that the maximum speed is taken into account for the cases where the critical speed was greater than the allowable speed. 28 vol. 36/2022 running safety factor of railway bridges using tail modelling p p v d pp p p dba s slab-subgrade kss & css slab track l krs & crs rail pad mb ,eib ,zb rail ms ,ei s ,zs mr ,eir ,zr figure 3. 2d computational model of the ballastless railway bridge. vcr = dfj i j = 1, 2, 3, ... i = 1, 2, 3, ..., 1/2, 1/3, 1/4, ... (14) where, d is the characteristic length of the coach and fj is the jth frequency of the bridge. 4. reliability problem as mentioned earlier, the running safety is implicitly controlled by limiting the vertical acceleration of the bridge deck. therefore, the limit state function is formulated as eq. (15), where failure occurs for g(x) ≥ 1. g(x) = amax(x)(1 + χm ) alim (15) where amax(x) is the maximum vertical acceleration of the bridge deck for x realization of the basic random variables, alim is the maximum allowable vertical acceleration of the bridge deck (here considered as 10 m/s2), and χm is the model uncertainty. the basic random variables considered are listed in table 1, with those related to the bridge and train properties taken from [29–31]. note that the target bridges have spans of 10 m, 20 m, and 30 m, and the values given in table 1 are in the same order. the mean values of the cross-sectional moment of inertia and mass per length correspond to the minimum allowable values satisfying conventional deterministic design methods [29]. the nominal values of the other parameters are taken from [6, 7]; the assigned distribution function and standard deviations are based on recommendations from [32] and the authors’ engineering judgment. 5. safety factor estimation the objective of the article is achieved by first calculating the maximum vertical acceleration of the bridges for passing trains with critical speeds. in this context, the sensitivity of the estimated safety indices is first evaluated with respect to the number of samples considered (see figure. 4). it should be emphasized here that the analyzes for this task were performed only for trains with a speed of 240 km/h. as can be seen, the method almost converges when the number of samples exceeds 10,000; however, 20,000 samples are considered in this study. 0 1 2 3 # sample 104 3 3.2 3.4 3.6 3.8 4 figure 4. sensitivity to the number of samples. then, the statistical parameters (i.e., mean and standard deviation) of demand are estimated by calculating the 99% quantile of these parameters using the bootstrap method. then, the system’s probability of failure is calculated using the tail approximation method. this is done by calculating the median of the probabilities of failure resulting from the methods discussed. these methods include fitting gpd by the maximum likelihood method using the hasofer, boos, and minimum mse thresholds, fitting the sigmoid function by the regression method using the minimum mse threshold, betalt, betaqh, and lnbetaqt. some examples of such approaches obtained for the 30 m span bridges are shown in figure 5. then, eq. (16) is used to estimate the existing safety factor of the conventionally designed bridges. this equation is developed in [10] for linear limit-state 29 r. allahvirdizadeh, a. andersson, r. karoumi acta polytechnica ctu proceedings variable dist.∗ mean/min./scale std/max./shape truncation bridge ib (m4) moment of inertia n [0.4, 1.5, 5.9] 0.01µib mb (kg/m) mass per length n [19000, 29000, 39000] 0.03µmb ξb (%)† damping ratio ln max(1.5 + 0.07(20 − l), 1.5) 0.3 eb (gpa) modulus of elasticity n 29.7 3.56 train d (m) coach length u 17 28 dba (m) axle distance u 2 3.5 p (kn) axle load w 194.93 9.14 ≤ 120 rail ir (m4) moment of inertia n 2 × 61.1e-6 0.01µir mr (kg/m) mass per length n 7850 × 15.38e-3 0.03µmr er (gpa) modulus of elasticity n 205 5 slab track ts (m) thickness n 0.3 0.01 bs (m) width n 2.5 0.005 ρs (kg/m3) mass density n 2500 100 ξs (%)† material damping ln 2 0.3 rail pad krp (mn/m) stiffness n 2 × 22.5 0.01µkrp ξrp (%)† damping ratio ln 10 0.3 slab subgrade ksg (mn/m3) stiffness n 100 0.2µkrp ξsg (%)† damping ratio ln 2 0.3 model uncertainty χm (-) model uncertainty n 0 0.086 ∗ n , ln , u and w indicate normal, lognormal, uniform and weibull distributions, respectively. † parameters are in physical space. table 1. considered basic random variables. functions using the separation function to linearize its standard deviation. γ = 1 + θβc cova 1 + kacova (16) where θ = 0.75 ± 0.06 for 1/3 < σ/µ < 3 is the separation constant linearizing the standard deviation of the linear limit state function, i.e. σz = (σ2r + σ 2 s ) 1/2 changes to σz ≈ θ(σr + σs ) for limit state functions in the form of g(x) = r − s. also, βc is the safety index corresponding to the limit state under consideration, cova is the coefficient of variation of the demand (here the vertical acceleration of the bridge deck), and ka is the standardized characteristic value (95% quantile denoted here as ak) of the demand, which reads as ka = (ak − µa)/σa. next, the new safety factor is calculated using the same data, replacing the existing safety index with a target safety index, which is βt = 3.719 (i.e., pf,t = 10−4). the calculated values are given in table 2; where γexist and γnew indicate an approximation of the safety factor based on conventional design methodology and the potential values to be adopted in the future, respectively. it should be emphasized here that minimum allowable mass values have been assigned to the bridges considered; therefore, the values presented represent an approximate lower bound of the possible safety indices. as can be seen, the conventional design method does not result in a consistent safety factor; although an identical factor was used for the design of all bridges. this conclusion underscores the point that safety factors in the future should probably be a function of span length, rather than being constant. in addition, bridges with shorter spans are more susceptible to higher accelerations, resulting in a lower safety factor compared to longer bridges. considering the safety factor obtained for bridges with a span of 20 m, it is even possible that some scenarios are neoconservative. on the contrary, criteria other than running safety may dominate the design of bridges with larger spans, e.g., displacement-related criteria such as passenger comfort [30]. therefore, a very large safety factor is estimated as a consequence of 30 vol. 36/2022 running safety factor of railway bridges using tail modelling 0.15 0.2 0.25 0.3 g 0.9 0.92 0.94 0.96 0.98 1 1p f emprical cdf gpd-mle-boos (a) . gpd-mle-boos threshold 0.1 0.2 0.3 g 0 0.2 0.4 0.6 0.8 1 1p f emprical cdf sigmoid-reg-best u (b) . sigmoid-regression-minimum mse threshold 0.1 0.2 0.3 g 0 1 2 3 4 emprical beta-lt (c) . beta-lt 0.1 0.2 0.3 g 0 0.5 1 1.5 l n( ) emprical lnbeta-qt (d) . lnbeta-qt figure 5. example of fitted tail models for bridges with span length of 30 m. l (m) γexist γnew 10 2.91 1.18 20 1.29 1.37 30 6.46 1.17 table 2. existing and proposed safety factor for running safety of high-speed ballastless railway bridges. a very small probability of failure. considering this, the reliability of the estimated existing safety factor for bridges with a span of 30 m is questionable. on the other hand, the proposed new values for the safety factor appear to be more consistent with a conservative value of about 1.4. however, as mentioned above, the use of a constant safety factor for all bridges with different spans needs to be revised. 6. conclusions in this study, the reliability of the conventional safety factor for the running safety of ballastless high-speed railway bridges was investigated. it was found that despite the use of a constant safety factor for bridges with different spans, the resulting designs may not have consistent safety factors. therefore, it seems necessary to revise the current safety factor by proposing a set of new factors that depend on the span length of the bridge. in this regard, rough proposals based on an arbitrary target safety level were recommended, showing the potential of using smaller values in future design regulations. acknowledgements this project has received funding from the shift2rail joint undertaking under the european union’s horizon 2020 research and innovation programme under grant agreement no 101012456. 31 r. allahvirdizadeh, a. andersson, r. karoumi acta polytechnica ctu proceedings references [1] cen.en 14363. railway applications testing and simulation for the acceptance of running characteristics of railway vehicles running behaviour and stationary tests. european committee for standardization, 2016. 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[8] r. allahvirdizadeh, a. andersson, r. karoumi. reliability assessment of the dynamic behavior of high-speed railway bridges using first order reliability method. in eurodyn 2020, 11th international conference on structural dynamics, athens, 23-26 november, 2020., vol. 2, pp. 3438–3450. 2020. https://doi.org/10.47964/1120.9282.18654. [9] p.-l. liu, a. der kiureghian. multivariate distribution models with prescribed marginals and covariances. probabilistic engineering mechanics 1(2):105–112, 1986. [10] r. e. melchers, a. t. beck. structural reliability analysis and prediction. john wiley & sons, 2018. [11] j. d. sørensen. notes in structural reliability theory and risk analysis. department of civil engineering,aalborg university, 2011. [12] r. melchers. importance sampling in structural systems. structural safety 6(1):3–10, 1989. [13] s.-k. au, j. l. beck. estimation of small failure probabilities in high dimensions by subset simulation. probabilistic engineering mechanics 16(4):263–277, 2001. [14] b. sudret. meta-models for structural reliability and uncertainty quantification. arxiv preprint arxiv:12032062 2012. [15] r. allahvirdizadeh. reliability-based assessment and optimization of high-speed railway bridges. licentiate thesis, kth royal institute of technology, 2021. [16] m. a. maes, k. breitung. reliability-based tail estimation. in probabilistic structural mechanics: advances in structural reliability methods, pp. 335–346. springer, 1994. [17] j. rocha, a. henriques, r. calçada. safety assessment of a short span railway bridge for high-speed traffic using simulation techniques. engineering structures 40:141–154, 2012. https://doi.org/10.1016/j.engstruct.2012.02.024. [18] p. ramu, n. h. kim, r. t. haftka. multiple tail median approach for high reliability estimation. structural safety 32(2):124–137, 2010. https://doi.org/10.1016/j.strusafe.2009.09.002. [19] a. hasofer. non-parametric estimation of failure probabilities. in f. casciati, b. roberts (eds.), mathematical models for structural reliability analysis, chap. 4, pp. 195–226. crc press, 1996. [20] d. d. boos. using extreme value theory to estimate large percentiles. technometrics 26(1):33–39, 1984. [21] j. caers, m. a. maes. identifying tails, bounds and end-points of random variables. structural safety 20(1):1–23, 1998. https://doi.org/10.1016/s0167-4730(97)00036-2. [22] t. ishigami, t. homma. an importance quantification technique in uncertainty analysis for computer models. in : first international symposium on uncertainty modeling and analysis, pp. 398–403. ieee, 1990. [23] p. lou. finite element analysis for train–track–bridge interaction system. archive of applied mechanics 77(10):707–728, 2007. https://doi.org/10.1007/s00419-007-0122-4. [24] cen. en 1991-2, eurocode 1: actions on structures part 2: traffic loads on bridges. european committee for standardization, 2003. [25] l. frỳba. dynamics of railway bridges. thomas telford publishing, 1996. [26] p. salcher, c. adam, a. kuisle. a stochastic view on the effect of random rail irregularities on railway bridge vibrations. structure and infrastructure engineering 15(12):1649–1664, 2019. [27] erri d 214/rp 9. rail bridges for speeds > 200 km/h. european rail research institute 1999. [28] l. frỳba. a rough assessment of railway bridges for high speed trains. engineering structures 23(5):548–556, 2001. https://doi.org/10.1016/s0141-0296(00)00057-2. [29] a. andersson, r. allahvirdizadeh, a. zangeneh, et al. high-speed low cost bridges: background report for in2track2, 2021. [30] r. allahvirdizadeh, r. karoumi, a. andersson. surrogate-assisted versus subset simulation-based stochastic comparison between running safety and passenger comfort design criteria of high-speed railway bridges. in 31st european safety and reliability conference (esrel 2021). 2021. https://doi.org/10.3850/978-981-18-2016-8_577-cd. [31] r. allahvirdizadeh, a. andersson, r. karoumi. minimum design requirements of high-speed railway bridges using reliability-based optimization. unpublished manuscript 2021. [32] jcss. jcss probabilistic model code. joint committee on structural safety (jcss), 2001. 32 https://doi.org/10.1080/23248378.2018.1503975 https://doi.org/10.47964/1120.9282.18654 https://doi.org/10.1016/j.engstruct.2012.02.024 https://doi.org/10.1016/j.strusafe.2009.09.002 https://doi.org/10.1016/s0167-4730(97)00036-2 https://doi.org/10.1007/s00419-007-0122-4 https://doi.org/10.1016/s0141-0296(00)00057-2 https://doi.org/10.3850/978-981-18-2016-8_577-cd acta polytechnica ctu proceedings 36:25–32, 2022 1 introduction 2 reliability assessment using tail modelling 3 computational model 4 reliability problem 5 safety factor estimation 6 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0295 acta polytechnica ctu proceedings 38:295–301, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague optimization model of solar cooling system with latent heat storage lana migla∗, kristina lebedeva riga technical university, latvia, faculty of civil engineering, department of heat, gas and water technology, k, ı̄psalas street 6a/6b, riga, latvia, lv-1048 ∗ corresponding author: lana.migla@rtu.lv abstract. today, the fastest growing end-use in the buildings sector is cooling demand, notes the international energy agency (iea). in the last few years, the interest in solar cooling systems has been growing globally. solar cooling systems are the next solution to meet the growing demand for air conditioning. the literature analysis shows that solar cooling systems mainly use single-stage absorption chillers, with water as the working fluid, and only a small fraction of them use phase change materials to improve heat storage. based on previous studies on solar-assisted cooling systems, the use of phase change materials for thermal storage should be investigated. the paper will present a study of a solar cooling system with the optimized pcm thermal storage, with the aim of stabilizing the operation of heat storage, taking into account the volatility of solar energy, the impact of short-term operations and peak hours on the amount of heat produced. this means that the amount of electricity consumed to heat up the accumulation tank will be significantly reduced. the optimization model in simulation software will be performed to test solar cooling system with latent heat storage, with aim to investigate the efficiency of the developed latent energy storage and provide technical guidance for the implementation of such system in the practice. all indicators, including environmental impact and economic calculations, will be identified in order to determinate the specific systems for foresight market uptake. keywords: solar cooling, thermal energy storage, phase change materials. 1. introduction the indoor climate in buildings has a major impact on occupant productivity, especially during the hot summer period when large amounts of energy are used to cool rooms. in 2020, global demand for space cooling continued to grow, in part due to greater home cooling with more people spending more time at home. space cooling accounted for almost 16 % of the final electricity consumption of the building sector in 2020 (approximately 1 885 twh) [1]. as solar cooling systems (scs) use solar thermal energy as a source of energy, they have the potential to substantially reduce the proportion of conventional energy used to cool rooms and the negative impacts they generate. however, there is a fundamental shortcoming of solar energy, a mismatch between available energy and consumption, which makes it necessary to store the energy. phase change materials (pcms) are effective for storing heat in thermal energy storage because they can store not only thermal energy beyond temperature change as sensible heat storage but also through melting and solidification, as latent heat storage, which occurs over a narrow temperature range allowing 5 to 14 times more thermal energy to be stored than sensible thermal energy storage of the same size [2]. latent heat storage uses the phase change of matter between different aggregate states – gas/liquid, solid/gas, solid/liquid and solid/solid [3]. based on a literature analysis [4, 5], the scs working in a temperature range of 60–120 °c were determined to be most efficient, depending on the thermal heat supplier used in the chiller and the capacity of the cooler. it is therefore necessary to look in depth at the possibility of integrating the various pcms into the energy storage to stabilize its functioning, thereby increasing the cop (coefficient of performance) of solar thermal driven air-conditioning system. 2. methods 2.1. solar thermal cooling systems in solar cooling systems a mixture of ammonia, lithium bromide (libr) and water can be used in the absorption cycle [6]. a solar assisted absorption cooling system using a mixture of 60 % libr and 40 % water has been investigated and the performance of the system has been optimized using various thermodynamic analysis methods. among other sorption cooling systems, absorption cooling circuits are the mainstream in the context of high cop, in particular at temperatures above 80 °c [7]. however, when comparing absorption cycle cop (0.5 to 0.8) to traditional cooling systems with cop up to 3.0, the absorption cycle efficiency optimization becomes significant [8]. the generator of the chiller is warmed up using heat from solar thermal panels. the heat is collected and stored in a thermal storage tank, after which the required amount of heat is supplied to the generator. it is 295 https://doi.org/10.14311/app.2022.38.0295 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en lana migla, kristina lebedeva acta polytechnica ctu proceedings figure 1. scheme of solar cooling system. estimated that the system heat carrier could reach 85–90 °c using the solar thermal panels [9]. the heat supplied to the generator will be used to evaporate the water from the libr-water solution mixture returning from the absorber. the simulation model of solar cooling system was developed with absorption chiller nominal power −8 kw and maximal cooling capacity of 11 kw. heat is mainly produced by solar collectors and used for thermal driving chiller. in the presented case, heat output of solar thermal system is 15 kw [10]. the solar thermal system includes one cubic meter thermal energy storage tank with an additional 8 kw electric heater. solar thermal cooling system model has the following parts: 1 – solar collectors, 2 – thermal storage tank with integrated pcm encapsulate tubes and auxiliary heater, 3 – chiller (evaporator, condenser) with cooling tower, 4 – cold storage and 5 – consumer part. basic configuration of solar thermal absorption cooling system technologies is shown in figure 1. a simulation model was developed and validated to assess the impact of the pcm on the operation parameters of the solar assisted cooling system. during the day solar available energy is volatile. it also has impact on the temperature in the thermal storage. this is shown in the case study where an experimental solar pilot plant was developed to verify the pcm behavior under realistic conditions, in which following stratification, the pcm modules were integrated at the top of the hot water storage tank. in addition, pcm module is designed to use multiple cylinders, notes luisa f. cabeza et al. [11]. the scientist concludes that the introduction of the pcm module into thermal water storage tanks for dhw supply is highly prospective technology. this will allow hot water to be supplied for a longer period, even without an auxiliary source, or use smaller tanks for the same application. as shown in maher shehadi et al. study [9], stable temperature for heating the generator has great importance in order to raise absorption cooling systems at cop. it can be seen in the figure 2. the additional cop of solar cooling system shall be determined as follows: copthermal = useful power [kw] input power [kw] . (1) when using pcm in a thermal energy storage tank, the additional useful energy will be the energy of the pcm storage. the cop with additional energy from hot water heating will be calculated by following equation: copthermal,modif ied = co + sho + who + upfpcms aspfc , (2) where co chiller output [kw], sho space heating output [kw], who water heating output [kw], upfpcms usefull power from pcm store, aspfc available solar power from collector [kw]. 2.2. pcm selection criteria in selecting pcms to be applied as lhtes materials for thermal storage, there are some important aspects to consider. the main criteria [12] for selecting pcms are: • thermophysical properties: pcm, within the required temperature range, has high latent heat of fusion, high conductivity, high specific heat, cyclic stability, high density, low phase transition volume change, low vapor pressure at operating temperature, and consistent melting. • kinetic properties: high nucleation rate and high crystal growth rate (achieving optimal recovery from storage system). 296 vol. 38/2022 optimization model of solar cooling system with latent heat storage figure 2. cop of the absorption cycle depending on the generator temperature, accounting for various absorber temperatures [9]. pcm/parameter mgcl2 · 6h2o rt100 erythritol pcm melting point, t m [°c] 116.7 99 117.7 pcm heat of fusion, lp cm [kj kg−1] 168.6 168 339.8 specific heat of pcm, liquid, cp,l [j kg−1°c−1] 2.6 (120 °c) 2.4 2.8 specific heat of pcm, solid, cp,s [j kg−1°c−1] 1.4 (20 °c) 1.8 1.4 thermal conductivity of pcm, liquid, kl [w m−1°c−1] 0.6 (120 °c) 0.2 0.3 (140 °c) thermal conductivity of pcm, solid, ks [w m−1°c−1] 0.7 (20 °c) 0.2 07 (20 °c) density of pcm, liquid, rl [kg m−3] 1450 (120 °c) 770 (130 °c) 1300 (140 °c) density of pcm, solid, rs [kg m−3] 1570 (20 °c) 940 (20 °c) 1480 (20 °c) table 1. main parameters of the selected pcm. • chemical properties: fully reversible cycle, chemically stable without deterioration after multiple freeze/melt cycles, chemically compatible with encapsulating materials, non-corrosive, non-toxic, nonexplosive, non-flammable. • economical characteristics: cost effectiveness and commercial availability for large-scale practice. • environmental characteristics: environmentally friendly, non-polluting over a lifetime, and has the possibility of recycling. based on the above criteria, pcm – mgcl2 · 6h2o, paraffin rt100 and erythritol to enhance the efficiency of a libr-h2o solar cooling system was selected. similar findings are also in place for other researchers [13]. their main parameters of selected pcm [14–16] are summarized in the table 1. highest energy density for a specified mass of all pcms was calculated using following equation: q = m[cp,l(t − tm) + λ + cp,s(tm − tref )], (3) where m [kg] is the mass of material, cp [j kg−1°c−1] is a specific heat of material, l [kw] is the latent heat, and t [°c] is the temperature. the thermal conductivity λ is calculated from the measured thermal diffusivity a, density r and heat capacity cp. 2.3. storage tank design for most latent heat applications, methods to improve heat transfer are required, as majority of phase change materials have inappropriately low thermal conductivity [17, 18] resulting in slow charging and discharging rates. series of methods have been used to optimize the heat transfer properties of pcm. among the most widely used methods are the finned tubes [19], rings and bubble mixing techniques [20], the application of carbon fibre and multitubes [21]. most of the studies report significant improvements in the thermal conductivity of phase change materials but have not compared them to other improvement techniques. from this it can be concluded that the design of the thermal storage tank is in charge of efficiently charging/discharging the pcm as needed. shell-and-tube heat exchangers are the most widely used type of storage technique, with pcm on the shell side. for example, francis agyenim [22] conducted a study 297 lana migla, kristina lebedeva acta polytechnica ctu proceedings on a variety of pcm tube configurations, where the heat transfer improvement for three different modifications – circular fins, longitudinal fins and multitube systems was compared. the findings indicate that the use of a heat transfer mechanism increases the amount of the charge/discharge energy. the quantity of charged/discharged energy depends on the effectiveness of the heat transfer technique used in the pcm modules. equally important is that the multi-tube system has the highest stored energy, followed by the longitudinal and circular finned systems in that order. the longitudinally finned system with the highest discharge efficiency, followed by the multi-tube circular finned systems. as the author [22] points out, the multi-tube system was most effective for charging, but exhibited significant subcooling during discharging process. multitube systems will require a minimum container size, followed by a system with vertical and circular fins. the combination of longitudinal fins and multitubes in each configuration is the optimal combination to boost the pcm charge and discharge efficiency. cylindrical geometries are said to be the best performing storage systems with high efficiency and cylindrical containers with minimal dimensions [23]. based on effective heat capacity method, heat capacity is considered to be a function of temperature due to phase changes in the temperature interval between melting and solidification, so the calculations are done on pcm, using both – temperature and latent energy. when charging (pcm melting) – the hot water from the solar collector exchanges heat with the solid pcm and the melting of pcm starts. whereas when discharging (pcm solidification) – the cold heat carrier flows in the inner tube and the pcm modules exchange heat with it to initiate the solidifying process. based on the mentioned above, two experimental configurations of pcm modules were studied. in the first experiment, pcm is filled in the copper shell tube with an inner diameter of 146 mm and the heat carrier flows into the embedded 54 mm tube without a heat increase. this served as a starting point for evaluating the efficiency of the other advanced configurations. in the second experiment, circular fins were welded in the surface of the heat transfer tube. the amount of cumulative energy charged and discharged, heat lost to the environment and storage efficiency were calculated using mathematical model [22] which contains the following equations: • cumulative energy charged/discharged qchar [j]: qchar = t∑ 0 ṁcp,h(∆t )∆t, (4) where ∆t = t0 − ti for charged energy, ∆t = ti − t0 for discharged energy, ṁ mass flow rate [kg/s], ∆t change in time [s], ∆t change in temperature [°c], cp,h specific heat capacity of heat carrier [kj/kg°c]. • cumulative energy heat lost qloss [j]: qloss = t∑ 0 ti,ins − to,ins ln(ro,ins/ri,ins) 2rkinsl + 1 h0a0 ∆t, (5) where ri,ins inner radius [m], ro,ins outer radius [m], h0 heat transfer coefficient [w/(m2 °c)], a0 area [m2], l length [m], kins thermal conductivity of material [w/m°c], ti,ins inner temperature [°c], to,ins outer temperature [°c]. • charging efficiency ηchar : ηchar = qchar − qloss qchar . (6) • discharging efficiency ηdischar: ηdischar = qdischar qchar − qloss . (7) • storage efficiency based on energy charged: ηstore,input = ηchar × ηdischar . (8) the efficiency of the heat transfer mechanism used in the pcm determines the cumulative amount of energy charged/discharged. 3. results after equation (3), it is calculated that the energy storage capacity using erythritol is the highest. erythritol results are 20 % higher than magnesium chloride hexahydrate (mgcl2 · 6h2o), moreover 32 % higher than rt100. this coincides with other studies [13]. in addition, pure mgcl2 · 6h2o shows supercooling about 37 k, thus the study will continue and seek appropriate modifications. therefore, in the experiment erythritol was used for filling in tubes. the results of two experimental configurations of copper tubes filled with erythritol are shown in figure 3. the temperature curve produced by the experiment shows the temperature spread over the three different phases during the charge/discharge process of experimental tube filled with erythritol. in the first phase the energy was accumulated by the pcm and was used to raise the temperature of the erythritol till it nearly reached melting point temperature 117.7 °c. the temperature increased almost linearly with time. in the 298 vol. 38/2022 optimization model of solar cooling system with latent heat storage figure 3. temperature curves of two experimental tube configurations melting/solidification. next phase temperature curves tended to asymptote due to a rapid change of phases. in the third phase the energy was absorbed in the form of sensible heat and the temperature increased once again, only then the temperature curve started to fall. due to the solidification a subcooling process effect was observed. the circular finned system improved the temperature curve a little during the charging phase of sensible heat addition when conduction controlled the heat transfer, despite the increase in surface area of the circular finned system over the tube without a fin system. in the discharge phase, the tube with fins shows slightly better results, but this difference does not appear to be significant. the cumulative amounts of energy charged and discharged were calculated using equations (4)–(8) and summarized in table 2. the overall coefficient of performance of real solar cooling systems was calculated by equations (1) and (2), the copthermal close to 0.78, but copthermal,modif ied 0.81 throughout the year. 3.1. energy and economic analysis of the system based on the simulation results, it was concluded that the modified thermal storage tank maintains the temperature of the heat carrier required for the operation of the scs operation for a long period of time, so using the latent storage tank significantly reduces the energy generated by the auxiliary heater. given that, the excess heat is used for dhw needs. it should be added that the modified storage tank simulation results show substantial outcome for providing stable operation of the solar collector system and compensating for short-term variations in available solar energy (eg. for example, cloudy hours) [24]. with the pcm modified storage tank, the cop increased by an average of about 4 % compared to the cop of the scs non-modified storage tank. in addition, the auxiliary heater activation time for additional heating has been reduced by 21 % compared to scs with non-modified thermal storage tank. this is due to more stable thermal insulation/retention which would reduce energy consumption and will increase the lifetime of the scs. the specific economical audit software tool is used to run simulations to determine total energy consumption and co2 emissions over the service life of the system. the main phases of the simulation tool include material and manufacturing process, as well as transportation and end-of-life potential. the cost of capital of the system can be categorized among the individual components, with collectors and the chiller which accounts for most of the system cost (30–40 %). next most expensive part is the cooling tower, then the control costs, and finally the thermal storage unit (about 5 %). the low cost of thermal storage provides the potential to enhance total system performance at a very little additional capital cost. 4. conclusions research investigates the efficiency of the developed latent energy storage for solar cooling system to increase cop of such systems. the main indicators, including environmental impact and economic calculations, were identified. the economic impact of modified thermal storage on solar cooling performance depends on many factors, including the electricity charges used in traditional compressor-based cooling systems. the possibility of using pcm as a storage instead of water also depends on the system properties such as operating temperature set points. this study presents an analytical methodology that allows the selection of the optimal size of thermal storage as well as to quantify the benefits of this thermal storage in terms of additional solar fraction. this can be used to maximize the added economic value of the storage for different 299 lana migla, kristina lebedeva acta polytechnica ctu proceedings parameters/configuration tube tube with circular fins cumulative energy charged [kj] 9555.0 10501 cumulative energy lost [kj] 552.3 589.3 effective energy charge [kj] 8665 9935 cumulative energy discharge [kj] 4966 7591.1 storage efficiency (%) 62.2 75.2 table 2. energy charged and discharged and storage efficiency. solar cooling systems. it has been found that pcm can reduce the demand for an auxiliary heat source during peak loads and low radiation periods. acknowledgements this work has been supported by the european regional development fund within the activity 1.1.1.2 “postdoctoral research aid” of the specific aid objective 1.1.1 “to increase the research and innovative capacity of scientific institutions of latvia and the ability to attract external financing, investing in human resources and infrastructure” of the operational programme “growth and employment” «latent heat storage for sustainable solar cooling» no.1.1.1.2/viaa/4/20/661. references [1] internation energy agency report 2018. [2022-04-01]. https://www.iea.org/reports/the-future-ofcooling [2] r. millers, a. korjakins, a. lešinskis, a. borodinecs. cooling panel with integrated pcm layer: a verified simulation study. energies 13(21):5715, 2020. https://doi.org/10.3390/en13215715 [3] r. zeinelabdein, s. omer, g. gan. critical review of latent heat storage systems for free cooling in buildings. renewable and sustainable energy reviews 82:2843–2868, 2018. https://doi.org/10.1016/j.rser.2017.10.046 [4] m. m. a. khan, r. saidur, f. a. al-sulaiman. a review for phase change materials (pcms) in solar absorption refrigeration systems. renewable and sustainable energy reviews 76:105–137, 2017. https://doi.org/10.1016/j.rser.2017.03.070 [5] k. du, j. calautit, z. wang, et al. a review of the applications of phase change materials in cooling, heating and power generation in different temperature ranges. applied energy 220:242–273, 2018. https://doi.org/10.1016/j.apenergy.2018.03.005 [6] s. a. a. ameer, h. a. k. shahad. state of the art of solar absorption cooling technologies. international journal for research in applied science & engineering technology 5(3):84–98, 2017. https://doi.org/10.22214/ijraset.2017.3017 [7] r. sekret, m. turski. research on an adsorption cooling system supplied by solar energy. energy and buildings 51:15–20, 2012. https://doi.org/10.1016/j.enbuild.2012.04.008 [8] u. eicker, d. pietruschka. design and performance of solar powered absorption cooling systems in office buildings. energy and buildings 41(1):81–91, 2009. https://doi.org/10.1016/j.enbuild.2008.07.015 [9] m. shehadi. optimizing solar cooling systems. case studies in thermal engineering 21:100663, 2020. https://doi.org/10.1016/j.csite.2020.100663 [10] k. lebedeva, l. migla. latent thermal energy storage for solar driven cooling systems. in engineering for rural development, vol. 19, pp. 1134–1139. 2020. https://doi.org/10.22616/erdev.2020.19.tf273 [11] l. f. cabeza, m. ibáñez, c. solé, et al. experimentation with a water tank including a pcm module. solar energy materials and solar cells 90(9):1273–1282, 2006. https://doi.org/10.1016/j.solmat.2005.08.002 [12] k. faraj, m. khaled, j. faraj, et al. a review on phase change materials for thermal energy storage in buildings: heating and hybrid applications. journal of energy storage 33:101913, 2021. https://doi.org/10.1016/j.est.2020.101913 [13] l. hosseini. design and analysis of a solar assisted absorption, cooling system integrated with latent heat storage. master’s thesis, delft university of technology, 2011. 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https://doi.org/10.1016/0017-9310(81)90035-1 vol. 38/2022 optimization model of solar cooling system with latent heat storage [19] f. agyenim, p. eames, m. smyth. a comparison of heat transfer enhancement in a medium temperature thermal energy storage heat exchanger using fins. solar energy 83(9):1509–1520, 2009. https://doi.org/10.1016/j.solener.2009.04.007 [20] r. velraj, r. v. seeniraj, b. hafner, et al. heat transfer enhancement in a latent heat storage system. solar energy 65(3):171–180, 1999. https://doi.org/10.1016/s0038-092x(98)00128-5 [21] f. agyenim, p. eames, m. smyth. heat transfer enhancement in medium temperature thermal energy storage system using a multitube heat transfer array. renewable energy 35(1):198–207, 2010. https://doi.org/10.1016/j.renene.2009.03.010 [22] f. agyenim. the use of enhanced heat transfer phase change materials (pcm) to improve the coefficient of performance (cop) of solar powered libr/h2o absorption cooling systems. renewable energy 87:229–239, 2016. https://doi.org/10.1016/j.renene.2015.10.012 [23] m. e. zayed, j. zhao, w. li, et al. recent progress in phase change materials storage containers: geometries, design considerations and heat transfer improvement methods. journal of energy storage 30:101341, 2020. https://doi.org/10.1016/j.est.2020.101341 [24] y. allouche, s. varga, c. bouden, a. c. oliveira. dynamic simulation of an integrated solar-driven ejector based air conditioning system with pcm cold storage. applied energy 190:600–611, 2017. https://doi.org/10.1016/j.apenergy.2017.01.001 301 https://doi.org/10.1016/j.solener.2009.04.007 https://doi.org/10.1016/s0038-092x(98)00128-5 https://doi.org/10.1016/j.renene.2009.03.010 https://doi.org/10.1016/j.renene.2015.10.012 https://doi.org/10.1016/j.est.2020.101341 https://doi.org/10.1016/j.apenergy.2017.01.001 acta polytechnica ctu proceedings 38:295–301, 2022 1 introduction 2 methods 2.1 solar thermal cooling systems 2.2 pcm selection criteria 2.3 storage tank design 3 results 3.1 energy and economic analysis of the system 4 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0522 acta polytechnica ctu proceedings 38:522–526, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the new industrial revolution: finding life for the buildings left behind dawn bilobran 313 historic preservation, united states of america correspondence: dbilobran@313hp.com abstract. from large generating facilities to modest neighbourhood substations, public power structures are an exercise in dichotomy. captivating yet mysterious, designed with both powerful function and beauty in mind. quietly playing a role in the development of cities and supporting the activities of home worldwide, the magic of heritage power plants, pumping houses and substations is often hidden behind metal gates and pressed brick facades punctuated with oversized steel windows and carved decorative ornamentations. efforts to achieve global goals of carbon-neutrality paired with advancements in infrastructure, utility distribution and alternative energies now forces the reconsideration of many unique historic resources. brimming with astounding potential, power generation facilities present unique challenges that can be deterrents to redevelopment. adaptive reuse celebrates the contributions of those who designed, constructed and operated the architectural and engineering marvels that powered the world while deterring exceptional building materials from languishing in landfills. as an increasing number of sites are decommissioned how can they be positioned to power new experiences for generations to come? what redevelopment tools are available to incentivize the adaptive reuse of industrial heritage, specifically public utility architecture? how do government-led approaches to adaptive reuse differ? keywords: power plant, adaptive reuse, rehabilitation, heritage, industrial, decommissioning. 1. introduction worldwide goals to achieve carbon-neutrality and promote sustainable energy production are driving advancements in the generation and distribution of power. past reliance on non-renewable and polluting energies are being replaced by wind, solar and biofuels. concurrently, heritage power facilities are reaching a predictable end of life and face replacement by cleaner and more efficient technologies. as the energy sector repositions, the futures of many historic power and utility buildings are called into question. the industrial revolution ushered in a new era of production and growth. reliable delivery of power became a paramount necessity for manufacturing centres and burgeoning municipalities. power plants, substations, gasworks, and waterworks housed massive mechanical operations but also received sophisticated architectural treatments. like public institutions of the period such as post offices, libraries and schools, the design of power buildings emphasized function and proportion without abandoning aesthetic enhancements. for decommissioned heritage power buildings, adaptive reuse of these unique sites is the only alternative to demolition and relegation to landfills teeming with construction waste. distinctive architectural design, high grade materials and robust community affinity are all arguments for retaining and reusing historic generating facilities. the primary challenges to adaptive reuse of utility architecture are overcoming perceived obsolescence and financing. while funding is an inherent issue in historic building rehabilitations, the building scale and potential site contamination of power plants can be daunting adversaries to embracing preservation and reuse. government agencies must support project success by administering financial incentives while fostering partnerships created to activate sites. each are essential to promoting adaptive reuse of legacy power generation and service buildings that have reached end of service. 2. public utility architecture: why adaptive reuse? 2.1. the municipal plant: civic duty and beauty rapid expansion of manufacturing and population required cities around the world to provide massive amounts of power demanded by such unprecedented growth. engineering advancements enabled the manipulation of impressive amounts of water, gas, and heat to generate substantial outputs while intricate systems were installed to deliver power directly to the consumer. no longer reserved for only the largest cities or the elite, soon lamp-lit streets, electric streetcars, and new home comforts became the norm. the expansion of public utilities as a civic amenity resulted in the construction of captivating buildings dedicated 522 https://doi.org/10.14311/app.2022.38.0522 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 the new industrial revolution: finding life for the buildings . . . to powering the populace. world expositions, fairs and architectural design contests fostered an exchange of ideas and competitive atmosphere amongst nations and states and the resulting influence on architectural design included power structures. in the united states, the 1876 centennial celebration in philadelphia sparked a period of growth and a new commitment to following established principles of design when planning cities and buildings. the chicago exposition in 1893 further cemented this philosophy and became a launch point for the city beautiful movement, american architect daniel burnham’s response to the dirty and chaotic urban areas resulting from rapid industrial growth. burnham advocated that beautification and order of public buildings and spaces were necessary for a functioning city and resulted in a positive behavioural impact upon the masses [1]. european approaches to architecture had long emphasized aesthetics for civic projects. the paris exposition of 1900 paid special attention to public power innovations, with a palace to electricity featuring state of the art technology as well as the competitive construction of two artfully designed chimney stacks. the french entry to the design competition was a masterpiece, with large shield representing industry and decorative brick patterns terminating in impressively stepped caps and relief of half rosettes, showing “what france could do in the way of making art out of a brick pile” [2]. in germany, the design of utility buildings “were given architectural expression which does not belie their true nature” while still expressing an “appearance of quaintness and charm” [3]. inspired by expositions and planning movements that encouraged enhancement of public buildings and spaces, municipalities commissioned architects with portfolios of museums, libraries, and universities to partner with engineers on the design of new power facilities. styles such as beaux-arts, neoclassical, renaissance revival and later art deco and moderne, were used to elevate plants from modest undertaking to canvas for architectural expression. despite being unnecessary for the facility’s primary functions, the classical treatment and adornment of municipal power buildings was embraced. the orderliness of activities occurring inside the plant was replicated in the classical balance and proportion applied to exterior elevations, punctuated by rows of oversize steel windows to maximize daylight and ventilation. the deliberate application of decorative elements transformed utilitarian plants and substations into captivating works of art. architects designed new facilities with decorative tile, ornamental brickwork and attractive lighting. decorative embellishments occasionally played directly on the building function, with the integration of lightning bolts or patterns mimicking waves or the ombre effect of a chemical reaction. doorways were formalized with pediments and sculptural ornamentation and prominently featured the municipal crest or year of construction. publicly accessible spaces were handsomely trimmed in marble and stone and provided the same interior appointments one would expect visiting a post office or city hall. ample exterior lighting assured that even the night-time onlooker would be awed by the plant’s magnificence. large-scale energy generation and distribution required that power plants, substations, gasworks, and waterworks be constructed in nearly every town and state. as cities grew in population and prominence, pressure increased to construct more noteworthy structures and the application of classical techniques to public buildings enjoyed a resurgence. these facilities played a vital role in supporting metropolitan development, manufacturing, and activities of daily life. municipalities regarded their ability to provide reliable power as a source of civic pride and a way to attract new businesses and residents. privatelyheld utility companies considered power plants to be testaments of corporate success, necessary to build legitimacy and attract new investors. both regarded new facilities as investments worthy of careful planning and consideration, including architectural treatments. in 1916 james n. hatch, a consulting engineer to power plant construction in the united states advocated to the western society of engineers that artistic touches to a power plant should not be overlooked, as the stability and success of any enterprise is judged consciously or unconsciously by the appearance of its physical property [4]. power plants and substations were provided the same architectural considerations as other prominent municipal institutions. with nearly every opening of a new power facility, the plant was proclaimed as “world’s largest” in some capacity. the design, engineering and construction of new power plants became its own competition amongst towns and utility companies. records for highest output, gallons pumped, or new technologies were routinely eclipsed as more facilities were completed and competed for such illustrious distinctions. grand openings were celebrated with ribbon cuttings and lauded with newspaper articles detailing every aspect of the design. equipment suppliers placed advertisements featuring crisp images of brand-new turbines, boilers and engines and noting the impressive performance capacities of machinery. municipalities and private utility companies alike commemorated new additions with postcards and ephemera that artistically depicted the facility and its surroundings. power plants symbolised achievement and signalled a strong future. by refusing to underestimate the importance of architectural design when planning hydroelectric plants or electric substations, decision-makers achieved mechanically superior yet beautiful facilities. generation plants and distribution substations received a uniform architectural treatment favoured by the company or were influenced by the existing architectural styles of the area. california’s pacific gas and electric company (pg&e) placed great emphasis on the de523 dawn bilobran acta polytechnica ctu proceedings sign and engineering of its facilities, recognizing that power stations could be objects of beauty without diminishing the brute forces concealed inside. this approach resulted in a series of substations with stunning stone ornamentation that appear plucked from paris or rome but were instead tucked into san francisco neighbourhoods. other pg&e facilities throughout california are spanish renaissance, with red-tiled roofs and stucco plaster exterior walls, to compliment the prevalent use of that style throughout the region [5, 6]. existing architecture and popular materials routinely influenced design choices at new facilities, especially for distribution substations sited in neighbourhoods or prominent districts. in 1925 plans for a new municipal rapid sand filtration plant in washington d.c. were required to pass examination by the fine arts commission of the national capitol due to the water filtration facility’s prominent setting along a busy highway near the federal seat of government. after multiple rounds of scrutinization, the team selected a style for the complex that complemented the existing buildings in the district [7]. michigan’s detroit edison company’s substations took stylistic direction from current residential building stock in the neighbourhood, using reclaimed timbers to elevate the design of its tudor revival stations. the emphasis on architectural treatment and materials juxtaposed with massive power and force is an exercise in dichotomy and the primary reason historic power facilities are deserving of preservation and reuse. 2.2. material assets: leverage not landfill world leaders are committed to reducing waste and pollutants to promote sustainability and combat climate change. the european union is proposing carbon-neutrality by 2050, supplemented with a fastapproaching goal to drastically reduce emissions by 2030 [8]. the united states aims to cut its greenhouse gas pollution by 2030 [9]. to facilitate these goals, large coal and gas-fired energy generation plants must either convert to cleaner processes or face decommissioning. at the same time, advancements in energy generation and distribution call into question the futures of aging municipal utility facilities, once postcard-worthy beacons of pride. in the absence of a plan that includes preservation, the site either proceeds directly to demolition or languishes in vacancy. removing legacy power plants from service does not end their negative environmental impact if demolition and disposal is the only path forward. demolition undertakings produce an enormous amount of construction waste and risk the release of harmful substances into the environment. while some building materials such as asphalt, concrete and certain metals provide an economic incentive to recover and recycle, others will persist in landfills. a 2018 study by the united states environmental protection agency (epa) estimated that the united states generates 600 million tons of construction and demolition debris with demolition activities responsible for more than ninety percent. of this total, the majority of waste was relegated to landfills [10]. in the european union, construction and demolition debris is the largest single waste stream, responsible for 374 million tons in 2016 [11]. diverting construction waste to secondary recycling markets to lessen the environmental impact of demolition remains an industry priority but barriers remain to a uniform embrace of recycled construction materials. at legacy power plants, demolitions are weeks-long undertakings, often involving controlled explosives and transporting waste to specialised landfills. encapsulated hazardous materials are disturbed and disbursed into the air, ground and water. the massive amounts of building materials used to construct power generation and distribution facilities, including reinforced concrete, steel, and millions of bricks better serve the environment as a repurposed building rather than conveyed to landfills. adaptive reuse is an inherently sustainable solution for heritage utility buildings that are removed from service and deters valuable materials from contributing to an already overwhelming burden of construction waste. 3. overcoming challenges to reuse 3.1. opportunity not obstacle classification as industrial site need not be a death knell for the preservation and adaptive reuse of power and utility buildings. adaptive reuse must be presented early as a viable redevelopment option so that stakeholders understand the opportunities embodied in these unique locations and financial tools to assist with rehabilitation. assessing project feasibility, assembling financial incentives and marshalling local support are critical steps to position adaptive reuse projects for success. leveraging community affinity is exceedingly valuable when assembling support for saving heritage assets, particularly with utility owners and political representatives. strong citizen attachment to a utility facility can be a differentiating factor in the decision by community leaders and private developers to repurpose or demolish the asset. owners seeking to reactivate decommissioned utility facilities without transfer of ownership may enter into publicprivate partnerships or create long-term leases with non-profit organizations. these agreements can remove maintenance and operational burdens, often costly even for a decommissioned site, while enabling adaptive reuse to move forward. in seattle, washington a collaboration to preserve the 1906 georgetown steam plant exemplifies the potential of such municipally driven partnerships. decommissioned in 1977, the generating facility is owned by seattle city light, the municipal electric provider. it was named 524 vol. 38/2022 the new industrial revolution: finding life for the buildings . . . a national historic landmark in 1984, the highest distinction for a historic site bestowed by the national park service. after extended dormancy, the complex reopened in 2014 to great fanfare for guided tours and art-based installations. sustained public interest in the site led seattle city light to seek a partner to maintain and operate the shuttered steam plant. after a proposal process, the newly formed non-profit georgetown steam plant community development authority assumed fundraising, rehabilitation and maintenance responsibilities while continuing the public access and artistic spirit synonymous with the site. the group entered a twenty-five-year lease with plans to reinvigorate the steam plant as a historic, scientific, artistic, and cultural resource [12]. strategic partnerships can catalyse rehabilitation and reuse while removing obligations from local governments or utility providers who are unable to devote necessary resources to preserve historic facilities or undertake ambitious development projects. perceived challenges to the rehabilitation of power plants must instead be embraced during adaptive reuse undertakings. the embodied industrial grit and grime is a valuable commodity which distinguishes energy generation buildings from other architectural heritage. rather than removal, character-defining features and exceptional architectural elements should be preserved and showcased during reuse. multistory ceiling heights create dramatic vertically oriented spaces and equipment such as turbines and boilers can remain in situ as visual links to the building’s history. exposed catwalks, overhead rigging and pully systems create interesting visual features once cleaned and secured. large complexes that offer multiple buildings and auxiliary structures can be activated with a multitude of complementary uses. facilities sited along water should embrace this amenity in rehabilitation plans. the features which differentiate power buildings from other historic resources must not be viewed as deterrents but rather as fascinating elements worthy of a sustainable second act. 3.2. empowering reuse heritage building rehabilitation can be financially daunting and the cost to repurpose power structures fluctuates widely depending on exiting conditions and end use. to deter demolition, it is essential for government agencies to provide financial tools to offset the remediation and rehabilitation of historic power facilities. these initiatives can transform preservation projects from insurmountable to viable economic development undertakings. for maximum impact, stakeholders must be effectively informed and prepared to take advantage of such programs. in the united states, government programs play a large role in encouraging the redevelopment of historic resources. sites of significance designated in the national register of historic places can receive a twenty percent tax credit for eligible building rehabilitation expenditures. undertakings utilizing the credit must follow the secretary of the interior standards for rehabilitation to ensure character-defining features are retained during the project and provide for the modernization of mechanical and plumbing systems as well as accessibility and seismic performance. as of 2022 thirty-nine states administer individual historic tax credit programs that can be twinned with the federal incentive to increase the return. grants supporting adaptive reuse undertakings are also deployed at government and institutional levels but are largely reliant on budget allocations and highly competitive [13]. addressing potential site contamination is important to determining next steps during the adaptive reuse of public utility buildings. as different historical uses created varying degrees of environmental pollution, in some cases the facility may not present stakeholders with as significant of a burden an initially anticipated. for projects where sites are polluted, there are remediation techniques to clean contamination, but they are costly endeavours. to encourage reclamation of polluted industrial sites the u.s. government offers multiple incentive programs to assess, clean and sustainably reuse contaminated properties, also called brownfields. multiple programs and federal tax incentives are administered by the environmental protection agency to underwrite the elimination of contaminants and revitalization of eligible locations [14]. it is common for adaptive reuse projects at industrial sites such as power plants to utilize both the historic tax credit and brownfield programs to counteract the costs of rehabilitation. despite existing initiatives there remains a disconnect between decommissioned property owners recognizing adaptive reuse as an attainable alternative to demolition. government agencies must do more to incentivize the reuse of legacy power plant facilities, dissuade the kneejerk reaction to demolish and encourage thorough site remediation. 4. conclusion global commitments to carbon-neutrality and advancing energy technologies are relegating heritage power facilities to decommissioned status with demolitions occurring at unprecedented rates. these sites are massively undervalued in site planning and the perceived obsolescence of discharged power buildings must instead be viewed as an exceptional development opportunity. character-defining elements that received thoughtful consideration by architects and engineers during design must receive similar contemplation by stakeholders determining the fates of decommissioned facilities. to combat the existing disconnect between disposition planning and adaptive reuse, owners and stakeholders must be empowered to envision the site outside its industrial functions and be informed of best practices and financial incentives that support adaptive reuse undertakings. 525 dawn bilobran acta polytechnica ctu proceedings government support for the adaptive reuse of heritage power facilities is a critical component to project success. financial programs that incentivize remediation and rehabilitation are vital and affirm a commitment to environmental stewardship and preserving industrial heritage but must be buoyed with additional funds. proponents must advocate for the expansion of incentive programs with additional monies specifically directed toward the redevelopment of legacy power plants which present enchanting adaptive reuse potentials but often present with daunting remediation realities. changes to eligibility requirements and construction timelines for federal rehabilitation incentives may be necessary to encourage the phased reactivation of large complexes. collaborations between public and private sector stakeholders are critical to developing sustainable new uses for historic utility buildings. global affinity for power heritage and the inherent community attachment to power generating facilities is often overlooked, underestimated, and can make a significant impact during redevelopment conversations. as we boldly commit to sustainable energy generation and building practices, we must resist the pervasive belief that heritage power plants, gasworks and substations are obsolete and beyond any capable service. with the proper fuel and partnerships, these majestic monuments to power can continue serving generations to come. references [1] c. moore. daniel h. burnham: architect planner of cities. houghton mifflin, boston, massachusetts, 1921. [2] r. grimshaw. the paris exposition chimneys. steam engineering 10(10):289–291, 1900. [3] i. k. pond. modern german architecture. the brickbuilder 36(10):243–247, 1915. [4] r. c. miller. the force of energy: a business history of the detroit edison company. michigan state university press, east lansing, michigan, 1971. [5] f. c. ayars. architecture of utility structures. southwest contractor 16(8):10–11, 1915. [6] c. f. piatt. the new san mateo substation. pacific service magazine 5(1):9–12, 1913. [7] p. o. macqueen. new rapid sand filtration plant, washington d.c. – part ii. american city magazine 35(5):645–648, 1926. [8] european commission. 2030 climate target plan, brussels, belgium, 2020. [9] white house fact sheet. president biden sets 2030 greenhouse gas pollution reduction target, washington dc, 2021. [10] united states environmental protection agency. advancing sustainable materials management: 2018 fact sheet. epa 530-f-20-009, washington dc, 2020. [11] european environmental agency. construction and demolition waste: challenges and opportunities in a circular economy, copenhagen, denmark, 2020. briefing no. 14/2019. [12] n. macdonald. seattle city light leases georgetown steam plant to local non-profit. powerlines, seattle, washington, seattle city light, 2021. [13] national park service. federal tax incentives for rehabilitating historic buildings annual report for fiscal year 2021, washington dc, 2022. [14] united states environmental protection agency. 2021 brownfields federal programs guide, washington dc, 2021. epa 560-b-21-003. 526 acta polytechnica ctu proceedings 38:522–526, 2022 1 introduction 2 public utility architecture: why adaptive reuse? 2.1 the municipal plant: civic duty and beauty 2.2 material assets: leverage not landfill 3 overcoming challenges to reuse 3.1 opportunity not obstacle 3.2 empowering reuse 4 conclusion references my title https://doi.org/10.14311/app.2022.36.0127 acta polytechnica ctu proceedings 36:127–134, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague establishing model uncertainty of sls reinforced concrete crack models applied to load-induced cracking christina helen mcleod university of kwazulu-natal, school of engineering, discipline of civil engineering, mazisi kunene road, glenwood, durban 4041, south africa correspondence: mcleodc@ukzn.ac.za abstract. although some research has been performed, serviceability limit state (sls) concrete crack models have yet to be calibrated fully in probabilistic terms in structural design standards. this is partly due to the fact that the sls is generally not the critical limit state in structural design. however, in some specialist structures such as water retaining structures, the sls such as cracking is the limiting design criterion, specifically the crack width required to control leakage and thus requires a proper probabilistic analysis. in probabilistic crack models, the reliability of the crack model is determined by the performance function whereby the design limiting crack width is greater than the estimated crack width calculated using the appropriate design crack model. as cracking in concrete is a random mechanism with a high degree of variability and crack models tend to be at least semi-empirical with inherent uncertainty, model uncertainty of the crack model is significant and is applied in the reliability model as a random variable. a database was established of both short and long term cracking experimental data for the tension load case to quantify model uncertainty. however, the data for long term cracking is limited which meant that the model uncertainty in this case was not definitively established. this paper discusses the determination of model uncertainty for tension cracking, which could be extrapolated to other models where data is limited. keywords: concrete crack models, model uncertainty, reliability. 1. introduction serviceability limit state (sls) concrete crack models are treated nominally in probabilistic terms in structural design standards. this is partly due to the fact that slss are generally not the critical limit states. however, in some specialist structures such as water retaining structures, sls cracking is the limiting design criterion, specifically the crack width required to control leakage. in addition, crack models are mostly empirically or semi-empirically derived so have not been assessed to any extent in reliability terms. however, cracking in concrete is a random mechanism with a high degree of variability indicative of a significant level of model uncertainty which, in turn, influences the level of treatment of model uncertainty in probabilistic analyses. retief (2015) [1] and holický et al (2016) [2] outlined a framework to assess model uncertainty. the relative influence of model uncertainty can be divided into model uncertainty classes [1] classified by an increasing level of treatment, as having nominal effect, significant effect or dominating effect, with, respectively. when model uncertainty is at least significant, it is deemed to be justified to treat model uncertainty as a random variable in the probabilistic model. this contribution discusses the quantification of model uncertainty for shortand long-term tension cracking, which could be extrapolated to other models where data is limited and where sls governs. the selection of the theoretical model for the general probabilistic model (gpm) is also important as model uncertainty is specific to the chosen model. load-induced tension cracking is controlled in structural concrete design standards by limiting the crack width for a given loading, section and reinforcement configuration. the maximum crack width is considered at the tension surface of the cross section. in probabilistic assessments of sls crack models, the reliability of the crack model is determined by the performance function g, expressed as g = wlim − θ. wpredict where wlim is the design limiting crack width, θ is the model uncertainty as a random variable and wpredict is the estimated crack width calculated using the appropriate design crack prediction model. as model uncertainty is specific to the prediction model concerned, suitable load-induced crack models were investigated, applicable to structures where serviceability cracking governs the design, such as reinforced concrete water retaining structures (wrs). the south african industry utilises the now withdrawn british standard bs 8007 (1987) [3] and is at present updating the standard sans 10100-1 (2004) [4] for the design of reinforced concrete structures by adopting the corresponding eurocode [5], concurrently with the development of sans 10100-3 (draft) (2015) [6] for the design of wrs. mcleod (2013) [7] in a deterministic analysis of typical wrs wall configurations, showed that the eurocode crack model was 127 https://doi.org/10.14311/app.2022.36.0127 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en christina helen mcleod acta polytechnica ctu proceedings conservative considering the reinforcement ratio compared to bs 8007’ especially for the tension loading case (> 38 % increase), for a limiting crack width of 0,2 mm. this has significant economic consequences if en 1992 adopted in south africa. the release of the fib model code mc 2010 [8] included an update of the fundamental crack model of fib mc 90 on which the eurocode crack model is based and allows for long term effects. the load-induced crack models for pure tension of bs 8007 (1987) [3], bs en 1992-1 (2004) [5], fib mc 2010 [8] were thus considered and compared to the results obtained for flexural cracking, as reported in [9] and [10]. the results of this research are also of relevance to researchers developing eurocode-related crack models and are applicable to structures other than wrs where serviceability cracking has a significant level of importance in the design thereof, such as bridges. the results of this research will also be utilised in the probabilistic assessment of sls cracking in existing structures such as wrs and bridges. 2. model uncertainty treatment uncertainty exists in all models and requires quantification as far as possible for proper reliability assessment and may be divided into two main types of uncertainty [1]. inherent random variability (aleatory) exists in model parameters such as material properties and loads. as these parameters are treated as random variables in the reliability model, this inherent variability is measured by their coefficients of variation. epistemic uncertainty (that due to incomplete knowledge and statistical uncertainty) may be measured by a model factor treated as a random variable in the reliability model. this model factor is defined as the ratio of measured to predicted values, in this case, crack widths. this ratio, or model factor (mf), is then quantified in terms of its statistical parameters of mean, coefficient of variation (cov) and probability distribution function (pdf). model uncertainty is a measure of the performance of the prediction model where the mean of model factors indicates bias in the model. thus, a mf mean of 1 indicates an unbiased model. a mf mean greater than 1 represents underprediction; conversely, a mean less than 1 indicates an overprediction by the model. the variation of model factors is indicative of the level of uncertainty in the model a higher variation signifying a higher uncertainty. for a model to be suitable over a reasonable range of design applications, it should behave consistently in terms of model factor bias and uncertainty over the range of application. interdependence between model uncertainty and parameters results in drift in the mf mean and an inconsistent model. with a large degree of randomness in concrete crack mechanisms and multiple influencing variables, crack models have high degree of variability (cov well in excess of 0,1 which is the reference value for general structures). model uncertainty (θ) is thus a significant variable in the reliability crack model, as shown by mcleod (2019) [9], requiring full probabilistic treatment. some research (e.g. [11]) has been done in establishing model uncertainty as a model factor but mostly for flexural short term cracking. the mean and cov of parameters were reported but not the pdf of model uncertainty necessary for reliability assessment. 3. deterministic analysis of predicted crack width in determining the predicted crack width for the mf, typical configurations of wrs and thus critical load cases for cracking were investigated. load-induced pure tension cracking occurs in the wall of circular water retaining structures (wrs) due to hoop stresses in the horizontal plane induced by water load, which is considered as a quasi-permanent load by both eurocode and sans design standards. the main reinforcement is thus placed in the horizontal direction with any tension cracks running perpendicular to this reinforcement. as tension cracks in this case tend to be through the full cross section, this is a critical load case for water tightness. a brief summary of the chosen design standard formulations used to calculate the predicted crack width for the critical pure tension load case in a wrs follows. a more comprehensive discussion may be found in [9]. crack mechanism philosophies differ in the modelling of the development and transfer of the stresses between the steel and concrete around a crack in the section, and the resulting strain incompatibility. as cracking is a serviceability limit state, linear elastic stress-strain theory applies. the predicted maximum crack width corresponds to a 5 % probability of exceedance by most design code formulations [9],[10]. this may not probabilistically match the measured maximum crack width, and therefore contributes to model uncertainty. 3.1. bs 8007 crack model the bs 8007 crack model [3] is a no bond-slip empirical model. the maximum surface crack width, w, for tension is calculated using: w = 3 acr εm (1) crack spacing is assumed to be a function of acr , as the distance from the crack considered to the nearest longitudinal reinforcing bar. the mean strain at the surface, εm, is calculated from: εm = ε1 − ε2 (2) where ε1 is apparent steel strain (εs) at the surface and ε2 is the concrete tension-stiffening effect. the equations to calculate tension stiffening strain for tension were derived empirically, calibrated to the specified maximum crack width limit (wlim) as follows: 128 vol. 36/2022 sls reinforced concrete crack models uncertainty ε2 = 2 bt h 3 esas f or wlim = 0, 2 mm (3) ε2 = bt h es as f or wlim = 0, 1 mm (4) es is the steel modulus of elasticity, h is section depth, bt is the width of the section in tension and as is the area of the tension reinforcement. interpolation for other crack widths is not possible which limits the application of this crack model. 3.2. en 1992 crack model the en 1992 crack width design equation [5] for pure tension cracking is a semi-analytical bond-slip model, developed from the compatibility relationship for cracking and assuming the stabilised cracking phase has been reached, namely: wm = srm · εm (5) where wm is the mean crack width, srm is average crack spacing and εm the mean strain. this crack width prediction model is based on the more fundamental model of fib mc 90. the mean strain is: εm = εsm − εcm (6) where εsm is the mean strain in the reinforcement under loading calculated using linear elastic theory. the mean concrete strain, εcm, also known as tension stiffening, is calculated using: εcm = kt fct,ef f ρct,ef f (1 − αe ρpef f ) / es (7) where αe is the modular ratio es/ec, ρp,ef f is the effective reinforcement ratio between the reinforcement area and the effective area of concrete in tension, fct,ef f is the mean tensile strength of the concrete at the time of cracking, ec is the concrete modulus at the time of cracking, and kt is a factor dependent on the duration of load. a minimum limit of 0, 6 σs/es (where σs is the steel stress) is placed on the mean strain. the design (or "maximum") crack width is required rather than the mean width. this maximum crack width is related to the mean crack spacing by the equation: wk = (βw srm) εm (8) where (βw srm) is the maximum crack spacing (sr,max) and the factor βw (the ratio of sr,max to srm) has a value of 1,7 corresponding to 1, 64× standard deviations from the mean (normal distribution of crack widths). the en 1992 maximum crack spacing is: sr,max = k3 c + k1 k2 k4 φ/ρp,ef f (9) where φ is the bar diameter, c is the concrete cover to the longitudinal reinforcement and k1 is a coefficient taking into account of the reinforcement bond properties, having a value of 0,8. the values of the factors k3 and k4 are determined by european union individual member countries’ national annexes, with recommended values of 3,4 and 0,425 given for k3 and k4, respectively. the factor k2 allows for stress distribution and has a value of 1 for pure tension loadinduced cracking. all k-values are empirical factors, therefore would contribute to model uncertainty. 3.3. mc 2010 crack model the fib mc 2010 crack model [8] is an update of the fib mc 1990 model. the design crack width, wd, as a maximum (assumed to be the 95th percentile, as en 1992) is determined from: wd = 2 ls,max(εsm − εcm − η εsh) (10) where ls,max is the transfer length over which slip occurs (equal to half the crack spacing) and εsh is the free shrinkage strain over time. the factor η is zero for short-term cracking, and 1,0 for long-term cracking. the minimum limit on mean strain of en 1992 is also specified by mc 2010. the transfer length is determined using: ls,max = k · c + 0.25 fctm τcms · φs ρp,ef f (11) where k is an empirical parameter to account for the influence of the concrete cover (k = 1, 0 can be assumed), τcms is the mean bond strength between steel and concrete (considered to be evenly distributed between two cracks) and φs is the nominal diameter of reinforcing bars. the ratio between the concrete tensile strength and mean bond strength (fctm/τcms) is 1/1,8 for stabilised cracking for both short and long-term loads. mc 2010 allows for crack width determination over both the crack formation stage and the stabilised cracking phase. the mc 2010 crack model is valid for c ≤ 75 mm. 4. quantification of crack model uncertainty the model factor for the chosen concrete crack models was defined as the ratio between measured crack width and predicted crack width, wexp/wpred. a database of experimental values for load-induced tension cracking was thus compiled to establish the stochastic parameters of model uncertainty for the crack models of en 1992-1-1, mc 2010 and bs 8007. the database was that from mcleod (2019) [9] updated to include more recent testing by gribniak et al (2020) [13]. the statistical parameters and probability distribution necessary for probabilistic analyses were determined for the model factors of each crack model using standard statistical test methods to a 95 % confidence level. the same procedure of analysis was followed 129 christina helen mcleod acta polytechnica ctu proceedings researcher element type testduration no. of samples** farra & jaccoud [12] ties with single reinforcing bar, square cross section. short-term 71 gribniak [13] ties with 4 no. reinforcing bars, cover varied. 10 samples total, repeat samples. short-term 4 hartl (1977), upm data [14], [15] ties with single or 2 no. reinforcing bars, squarecross section. short-term 48 hwang [16] slab elements reinforced in both directions, axial tension in one direction. variation of cover and reinforcement short-term 34 wu [17] ties with single reinforcing bar, square cross section. short (7) &long-term (4) 7 + 4 eckfeldt [15] ties with 1or 2 no. reinforcing bars, square cross section. repeats 2 × 4 no. ties & 3 × 1 ties long-term 11 *final load steps considered only, en 1992 minimum strain complied with. table 1. sources of experimental data direct tension load-induced cracking. as detailed in [9] and [10]. non-parametric normality tests such as kolmogorov-smirnov (using lilliefors significance correction) and shapiro-wilks (corrected), as appropriate, were performed to a significance p of 0,05 to establish the estimated probability distribution of the model factor. graphical methods such as probability plots and box plots were used to validate the estimated probability distribution. through the statistical analysis of model factors, a first assessment of the crack models’ performances was done. pearson’s correlations were used to investigate the relationship between significant parameters and model uncertainty, and thus assess the consistency of each crack model. 4.1. experimental database with data for existing wrs lacking and challenging to obtain, experimental data was used to establish model uncertainty of the crack models of en 1992, bs8007 and mc2010. both shortand long-term data was considered, given the quasi-permanent nature of the water load in wrs. however, the data for long term tension cracking [9] is limited. sources for the database are summarised in table 1. given that the first filling of a wrs only occurs once concrete has reached at least its 28-day strength and considering the quasi-permanent nature of the water load, only data from the stabilised cracking stage was considered. where steel stress is small, crack widths may be underpredicted resulting in an overestimation of model uncertainty, unduly increasing the upper tail of the model uncertainty distribution. in addition, the applied loads and section geometry of water retaining structures typically result in a mean strain well above the minimum strain limit. therefore, data was selected where the calculated mean strain was at least the specified minimum limit of 0, 6σs of en 1992 and mc 2010 where σs is the steel stress determined using linear elastic theory. crack widths measured on the final load step only were considered to ensure independence of samples. results for any repeat samples were averaged to prevent undue sample bias. analyses were performed for en 1992 and mc 2010 using the updated database, whilst the results for bs 8007 were taken from [9]. 5. discussion on model uncertainty quantification 5.1. short term tension the final sample size for short-term tension was 86 after repeats were averaged, which is sufficient to obtain a reasonable estimate of model uncertainty. the statistical parameters for model uncertainty are summarized in table 2. the values obtained for the statistical parameters are similar to those reported on in [9]. there were some small differences on including the research of [13] which consisted of ties reinforced with multiple bars rather than the mostly single bars or uniform samples of the database of [9]. the statistical analysis showed that the en 1992 crack model tends to be conservative, overestimating the predicted short-term tension crack widths, with a mf mean of 0,75. the mf mean of around 1 obtained for the mc 2010 crack model shows little bias in this model. the mc 2010 mf cov of 0,32, however, is higher than that of the en 1992 model (cov of 0,25). the bs 8007 crack model underpredicts crack widths with mf means over 1,27. this bias increases significantly when tension stiffening is determined using the smaller limiting crack width, which is problematic when developing a gpm as the model is not consistent over crack widths away from the given limiting crack width (either 0,1 or 0,2 mm). there is also some uncertainty in the determination of acr in ties reinforced with a single bar as the formulation for acr results in underprediction of the crack spacing and width. 130 vol. 36/2022 sls reinforced concrete crack models uncertainty statistical parameter en 1992 mc 2010 bs 8007 wlim 0.2 mm [9] bs 8007 wlim 0.1 mm [9] mean 0.747 0.996 1.271 1.430 standard error 0.020 0.034 0.032 0.041 median 0.722 0.956 1.225 1.398 standard deviation 0.183 0.319 0.290 0.369 cov 0.245 0.321 0.228 0.258 sample variance 0.033 0.102 0.084 0.136 kurtosis 0.051 -0.581 -0.056 1.234 skewness 0.540 0.427 0.441 0.777 range 0.927 1.408 1.516 2.139 minimum 0.374 0.416 0.582 0.657 maximum 1.301 1.824 2.097 2.796 pdf ln ln n n count 86 86 82 82 table 2. model uncertainty statistical parameters short-term tension cracking. figure 1. probability plots for en 1992 and mc 2010 short-term tension cracking. referring to table 2, both the en 1992 and mc 2010 crack models exhibit a positive skewness which suggests that the distribution is not normal. for the en 1992 model uncertainty, the shapiro-wilks nonparametric test rejected the null hypothesis, whilst for the mc 2010 crack model, the shapiro-wilks significance factor was just greater than 0,05. thus, it is estimated that both models have a non-normal distribution for model uncertainty. curve-fitting was done for both crack models, considering normal and lognormal distributions. this indicated that both en 1992 and mc 2010 tends towards a lognormal distribution, as shown in figure 1. considering that a lognormal distribution produces lower reliability estimates than a normal distribution (so more conservative), and that the distribution of the mc 2010 crack model has a positive skewness and is not clearly normal, a lognormal distribution is assumed for both models. bs 8007 model uncertainty appears to have a normal distribution. 5.2. long term tension statistical analyses for long-term tension cracking are summarised in table 3, and as reported in [9]. the sample size given is the final one after all repeats were averaged. comparing the long-term model uncertainty statistical parameters [9] to the short-term case, crack models that do not take long term shrinkage into account (bs 8007 and en 1992) do not have a consistent model uncertainty as load duration increases. the en 1992 model uncertainty mean increases from 0,75 to 0,90 for short to long-term tension loading. however, when considering the mc 2010 crack model, the mf means for short and long-term loading are very similar. the cov for all models for long-term loading is lower than that of short-term tension, but still demonstrate the significant influence that model uncertainty has on the crack models. however, some restraint is required on interpreting the data given the small sample size. the sample size is too small to give a good indication of skewness, nevertheless, 131 christina helen mcleod acta polytechnica ctu proceedings statistical parameter en 1992 mc 2010 bs 8007 w 0,2 mm bs 8007 w 0,1 mm mean 0.895 0.988 1.318 1.603 standard error 0.078 0.076 0.211 0.280 median 0.860 0.946 1.089 1.353 standard deviation 0.220 0.214 0.597 0.793 cov 0.246 0.216 0.453 0.495 sample variance 0.048 0.046 0.356 0.629 kurtosis -1.539 -0.059 3.779 2.151 skewness 0.442 0.806 1.882 1.447 range 0.571 0.618 1.807 2.328 minimum 0.656 0.764 0.836 0.935 maximum 1.227 1.382 2.643 3.262 pdf (estimated) ln ln n n count 8 8 8 8 table 3. model uncertainty statistical parameters for long-term tension cracking [9]. a lognormal rather than normal distribution is suggested from the positive skewness values obtained for both the en 1992 and mc 2010 crack models. as found for short-term tension, the bs 8007 model uncertainty statistical parameters are not consistent for the different tension stiffening models. 5.3. evaluation of crack models pearson’s correlations the interdependence between selected parameters and model uncertainty for each model were evaluated, together with further assessment of the performance of the en 1992, mc 2010 and bs 8007 crack models by means of scatter plots, regression analyses and pearson’s correlations. the selected parameters were reinforcing ratio (as % as), steel stress, concrete tensile strength (fctm), section depth (h), concrete cover (c) and section width (b). the pearson’s correlation factors between the model factor and selected parameters are summarised in table 4 for each load case and crack model, respectively. the small sample size means that the long-term tension cracking correlations of [9] need to be viewed with caution. little variation in the sample configurations for long-term tension loading [9] also results in correlations relating to section geometry being overestimated for all models. this applies to some extent to short-term tension loading, as although the sample size is larger, many samples were limited to ties with a single reinforcing bar or have little variation. adding the experimental programme of [13] (which consisted of repeat samples with multiple bars as opposed to much of the more uniform database of [9] with ties reinforced with single bars) resulted in differences to the short-term tension values obtained for the pearson’s analyses reported in [9] for mc 2010 and en 1992 even though the sample size increase was small. further research using different configurations/real structures is required for tension cracking for a more accurate evaluation of interdependence between these parameters and the model factor. referring to table 4 and short-term tension, with the exception of the section depth to width ratio (h/b) for mc 2010, model uncertainty for the en 1992 crack model has low to moderate correlations with all parameters. as would be expected, model uncertainty has a negative correlation with concrete tensile strength (fctm) for both en 1992 and mc 2010. the inconsistency in the bs8007 crack model is clearly demonstrated as the pearson’s correlations vary substantially between the tension stiffening formulations. a larger sample size is required for a true reflection of the pearson correlations for the long-term case [9], as can be evidenced by the extremely varied values between shortand long-term loading. 5.4. comparison to flexure model uncertainty as the database for long-term tension cracking was very limited, comparisons were made to analyses to quantify model uncertainty as reported in [9] and [10] for shortand long-term flexure loading applied to the en 1992, mc 2010 and bs 8007 crack models. the statistical parameters for the mf for flexural cracking are summarised in table ??. further detail on the flexural crack models analyses can be found in [9, 10]. comparing the short-term model uncertainty for tension to that of flexure, the means for the mc 2010 crack model are similar at approximately 1, although the variation for the tension case is less than that of flexure. this is probably partly due to the greater uniformity of sample configurations of the tension database compared to flexure. lognormal distributions are indicated for the en 1992 and mc 2010 crack models for both flexural and tension cracking. however, model uncertainty parameters are similar enough to conclude that it is not necessary to distinguish between the flexural and tension load cases if the mc 2010 crack model is considered for the gpm. although the dataset for long term flexural cracking 132 vol. 36/2022 sls reinforced concrete crack models uncertainty load case parameter correlation with model factor en 1992 mc 2010 bs8007 w = 0, 2 mm [9] bs8007 w = 0, 1 mm [9] short-term tension steel stress 0.052 0.204 -0.284 -0.364 h/b 0.579 0.722 0.049 0.206 c 0.286 0.420 0.332 0.230 bar dia, φ 0.518 0.570 -0.068 -0.241 fctm -0.294 -0.503 -0.030 0.080 % as 0.352 0.390 -0.036 -0.019 long-term tension [9] steel stress 0.825 0.641 0.840 0.777 h/b 0.490 0.335 0.602 0.479 c -0.825 -0.641 -0.840 -0.777 bar dia, φ -0.414 -0.209 -0.544 -0.407 fctm -0.826 -0.724 -0.792 -0.773 % as -0.349 -0.156 -0.426 -0.334 table 4. pearson’s correlation matrix between model uncertainty & model parameters. load case statistical parameter en 1992 mc 2010 bs 8007 w = 0.2 mm bs 8007 w = 0.1 mm short term mean 1.107 1.052 1.185 1.112 cov 0.397 0.376 0.380 0.459 pdf ln ln ln ln count 164 164 164 164 long term mean 1.443 1.127 1.502 1.514 cov 0.331 0.380 0.336 0.357 pdf ln ln ln ln count 30 30 30 30 table 5. summary of statistical parameters of mf for flexural cracking [10]. [9, 10] is larger than that of the tension load case, further research for long-term loading in general is recommended, including expanding the database and application to real conditions, for a more accurate estimation of the statistical parameters of model uncertainty. pearson’s correlations were performed for flexural cracking, as reported by [10]. low to moderate correlations were found for all parameters except for % as for long-term flexure (r of 0,58). it is noted that the dataset for flexural cracking had more variation in terms of section and reinforcement configurations than that of tension cracking, which supports the suggestion that the uniformity of the dataset for tension cracking results in apparent dependence between these parameters and the mf. 6. summary and concluding remarks in quantifying model uncertainty for pure tension crack models, the significant degree of uncertainty was confirmed, which justifies a full probabilistic treatment of model uncertainty and sls crack models. the model uncertainty statistical parameters were determined for short and long-term tension cracking, although the latter case does require further investigation to confirm values obtained. however, considering the mf statistical parameters determined for short and long-term flexural cracking ([9, 10]), the shortterm mf values for tension cracking may be used for long-term mf’s when considering crack models such as mc 2010 which take long-term shrinkage into account. indications are also that for the mf, it is not necessary to distinguish between flexural and tension load-induced cracking. this simplifies the reliability assessment of load-induced cracking and the development of the gpm. in terms of crack model performance, the bs 8007 crack model is shown to be inconsistent as its tension stiffening model is dependent on the specified limiting crack width. any variation in crack width away from this limit results in a high variation in the mf, and thus an inconsistent model. this is a point to note for any empirically-developed model which may be considered for a gpm. it was also noted during the analyses that the point on the cross section where the crack was measured was generally not given in the experimental 133 christina helen mcleod acta polytechnica ctu proceedings records, introducing some further uncertainty into the predicted crack widths due to potential mismatches between measured and estimated crack positions. using acr based on the geometry of the cross section, rather than a crack spacing formulation, for ties with single bars resulted in underprediction of the measured crack widths. this effect would be mitigated in structures under real conditions. crack models that omit shrinkage strain (en 1992 and bs 8007) underpredict crack widths in the long-term loading case. mc 2010 was the most consistent crack prediction model and will thus be the basis for the gpm. further to this research, data will be compiled for existing structures such as bridges and water retaining structures for use in the probabilistic analysis of the gpm. the experimental database was compiled and refined with wrs in mind, where sls cracking is dominant in design, so focused on small crack widths. for long-term tension cracking in particular, it must be noted that as many of the elements tested were mostly ties reinforced with single reinforcing bars, the crack widths recorded may not necessarily be representative of real conditions in structures such as wrs where there are obviously multiple reinforcing bars in both the main and transverse directions. this was noted in [9] and [10], borne out in the further analyses. this requires further research to properly assess the influence of section and reinforcement configurations on tension cracking, and which will include data from existing structures. references [1] j. v. retief. contributions to the implementation of the principles of reliability to the standardised basis of structural design. dr of eng thesis, stellenbosch university, south africa. supervisor gap van zijl, co-supervisor c viljoen, 2015. [2] m. holický, j. v. retief, m. sýkora. assessment of model uncertainties for structural resistance. probabilistic engineering mechanics 45:188-97, 2016. https: //doi.org/10.1016/j.probengmech.2015.09.008. [3] bs 8007. design of concrete structures for retaining aqueous liquids. british standards institute, 1987. [4] sans 10100 -1 (2004): the structural use of concrete: part 1: design. south african standards division. pretoria, 2004. [5] bs en 1992-1-1: 2004. eurocode 2: design of concrete structures part 1-1: general rules and rules for buildings. british standards institute, 2004. [6] sans 10100 part 3 (draft): design of concrete liquid retaining structures. sabs-tc98-sc2 working group on implementation of south african national standard for liquid retaining structures. south african standards division, pretoria., 2015. [7] c. h. mcleod. investigation into cracking in reinforced concrete water-retaining structures. msc eng thesis, stellenbosch university, south africa. supervisors profs ja wium & jv retief, 2013. [8] g. l. balázs, p. bisch, a. borosnyói, et al. design for sls according tofibmodel code 2010. structural concrete 14(2):99-123, 2013. https://doi.org/10.1002/suco.201200060. [9] c. h. mcleod. model uncertainty in the prediction of crack widths in reinforced concrete structures and reliability implications. phd thesis, stellenbosch university, south africa. supervisor prof c viljoen, 2019. [10] c. h. mcleod, c. viljoen. quantification of crack prediction models in reinforced concrete under flexural loading. structural concrete 20(6):2096-108, 2019. https://doi.org/10.1002/suco.201900036. [11] m. lapi, m. orlando, p. spinelli. a review of literature and code formulations for cracking in r/c members. structural concrete 19(5):1481-503, 2018. https://doi.org/10.1002/suco.201700248. [12] b. farra, j.-p. jaccoud. influence du béton et de larmature sur la fissuration des structures et béton. rapport des essais: de tirants sous deformation impose de courte durée. institut de statique et structures beton armé et preconstraint. publication no. 140, 1993. [13] v. gribniak, a. rimkus, a. pérez caldentey, et al. cracking of concrete prisms reinforced with multiple bars in tension-the cover effect. engineering structures 220, 2020. https: //doi.org/10.1016/j.engstruct.2020.110979. [14] a. pérez caldentey. upm experimental data. personal communication via email, 2016. [15] l. eckfeldt. schlussbeicht verbesserug de vorhersagequalitat um sehr kleinen vissbreiten. phd thesis, tu dresden, germany. supervisor prof ehm curbach, 2009. [16] l. s. hwang. behaviour of reinforced concrete in tension at post-cracking range. msceng (civil) thesis. university of manitoba, canada, 1983. [17] m. h. q. wu. tension stiffening in reinforced concrete: instantaneous and time-dependent behaviour. phd thesis, university of new south wales, australia. supervisor prof ri gilbert, 2010. 134 https://doi.org/10.1016/j.probengmech.2015.09.008 https://doi.org/10.1016/j.probengmech.2015.09.008 https://doi.org/10.1002/suco.201200060 https://doi.org/10.1002/suco.201900036 https://doi.org/10.1002/suco.201700248 https://doi.org/10.1016/j.engstruct.2020.110979 https://doi.org/10.1016/j.engstruct.2020.110979 acta polytechnica ctu proceedings 36:127–134, 2022 1 introduction 2 model uncertainty treatment 3 deterministic analysis of predicted crack width 3.1 bs 8007 crack model 3.2 en 1992 crack model 3.3 mc 2010 crack model 4 quantification of crack model uncertainty 4.1 experimental database 5 discussion on model uncertainty quantification 5.1 short term tension 5.2 long term tension 5.3 evaluation of crack models pearson's correlations 5.4 comparison to flexure model uncertainty 6 summary and concluding remarks references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0111 acta polytechnica ctu proceedings 40:111–116, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague generation of ldpm structure formed by voronoi cells jan vozáb∗, jan vorel czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jan.vozab@fsv.cvut.cz abstract. a preliminary study of an approach to internal structure generation used in lattice discrete particle models (ldpms) [1]. the presented method used for particle generation and placement is intended to help realistically capture the internal structure of materials. first, a method for structure generation using ldpm is presented. then, the method of particle generation using a voronoi diagram [2] is described. the last part is the optimizations on the algorithm that use apollonius circles to calculate the specific points of the voronoi diagram. keywords: ldpm, voronoi diagram, voronoi cell, internal structure. 1. introduction the use of composite materials such as reactoplasts is increasingly common in today’s industry. the construction industry is no exception. however, other industries such as aerospace or automotive differ in several important ways. this mainly concerns the degree of material curing at which these materials begin to load. this effect has a strong influence on the response of the material under load, especially in combination with viscoelasticity. at the same time, the material properties change over time. one of the structural elements used in construction where these materials are involved involves chemical anchors. particle-filled polymers act as a mortar binder layer between the anchor and the building element to which they are attached. these mortars are considered as homogeneous materials. but at lower scales, they can be characterized as heterogeneous materials [3]. in [4], is described that the modification of conventional continuum models (phase-field and peridynamic) is not possible without discretizing the internal structure. for this reason we decided to use a discrete model. figure 1 shows the real internal structure of such a polymer mortar. the polymer binder that connects the grains is shown in gray and shown in light color. it is clearly shown that the grains are of different sizes. the failure of these materials is mainly around the grains in the mortar [5, 6]. breakage through the grains occurs mainly in high-strength concretes [1]. this allows us to use the lattice discrete particle model (ldpm), which creates the internal structure of the material. it is convenient for the materials we are investigating. the ldpm material model response is dependent on particle distribution, and thus multiple simulations are needed to provide credible results [7]. however, as can be seen in the figure 1, most of the grains are not circular. the ldpm model only allows the generation of circular (2d) or spherical (3d) particles. our goal is to modify the ldpm model to allow the generation of grains of different sizes and shapes. therefore, we decided to modify the particle generation using a voronoi diagram. the computational complexity of this model is also an important factor. this paper focuses on the modification of the algorithm and its subsequent optimization. figure 1. detail of internal structure: hilti hit re 500. 2. ldpm structure first of all, it is important to describe the original method of generating the mesosctucture using the ldpm method [1]. we are given the sample geometry and the prescriptive gradation curve, i.e., the maximum and minimum aggregate size, as input for structure generation. the maximum aggregate size defines the upper limit of the gradation curve. in contrast, the minimum aggregate size defines its lower boundary, i.e., the 111 https://doi.org/10.14311/app.2023.40.0111 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en jan vozáb, jan vorel acta polytechnica ctu proceedings figure 2. delaunay triangulation result given set of particles. figure 3. ldpm structure for circular particles. diameter below which no discretized particles are generated and located. thus, the minimum aggregate size affects the fineness of the discrete network and the computational cost. the process begins with the generation of particles on the exterior of the sample. in our case (2d), points are first generated at the corners and then in the middle of the sides of the square. in the figure 2 they are shown by the blue circular points. this is followed by particle generation on the interior of the sample. the particles are placed from largest to smallest in a random position to ensure that the sample contains as many as possible. in the figure, these particles are represented by circles of different colors. the nearest particles are then detected using delaunyan triangulation, which results in the underlying polytopes in the given space. they represent triangles in the plane and tetrahedrons in space. in the figure 2 they are shown figure 4. structure of circular voronoi particles before composition. with blue lines connecting the individual particles. the following process divides the triangle into parts that belong to the particles that form its vertices. the centres of the sides and the centre of the triangle are calculated, and the part of the triangle is connected to the particle to which it belongs. the individual segments are then combined to form the resulting polyhedral cells (figure 3). when all the polyhedral particles have been formed, their interaction with each other has to be determined through the facets formed by the splitting of the triangles. the displacements and rotations of these neighboring particles form the discrete compatibility equations in terms of rigid-body kinematics. on each facet of the cell, an inter-scale constitutive law is formulated to simulate cohesive fracture, compaction due to pore collapse, frictional slip, and velocity effect. finally, equilibrium equations are formulated for each individual particle. 3. voronoi structure this article studies the potential use of voronoi cell tessellation [2] for the ldpm model. the voronoi cells are particles with a boundary representing half the distance between the particles from which tessellation occurs [8]. boundaries in the fundamental case are represented by lines for points in 2d and planes for points in 3d. for different particles, the voronoi cells are more complex. the simplest planar particles in the plane are circles. when the radii are equal, the boundary is represented by a straight line. when the radii differ, the boundary becomes curved. considering more complex particles, such as ellipses, the boundary complexity becomes even greater. however, it is possible to generate a structure for particles of any shape. in the figure 4 you can see the cell structure created for the same set of particles as for the ldpm struc112 vol. 40/2023 generation of ldpm structure formed by voronoi cells figure 5. structure of circular voronoi particles after composition. ture. the challenging thing about forming a structure using voronoi cells is that if the particles have different radii, or if they are not circular, the boundary between them is formed by a curve. in the figure you can see that in particular, the boundaries of cells formed by particles on the perimeter of a square are curves. furthermore, this effect can be seen for particles that are close together. we deal with this effect by generating more points on the boundary of the particle, thus generating a special cell. once the entire structure is generated, the points belonging to each particle are merged into a complete cell that takes into account the size and shape of the particle. the resulting boundary is a semilinear curve. result is shown in the figure 5. in this case, the particles are represented by 50 points on the particle boundary. for the calculation we used the matlab software [9]. compared to the structure formed by the ldpm method, they are smoother. smaller particles are connected to fewer other particles. at the same time, there are no fragments like those seen in the upper left corner of the picture. for the purple particle, the cell boundary passes through the particle itself. the figure 6 shows a structure made up of elliptical cells. at first glance, it is obvious that the boundaries between the particles are much more curved due to the complexity of the particle shape. therefore, even when the particles are represented by 44 points, small fluctuations in the smoothness of the boundary can be seen. nevertheless, in the figure 7 the boundaries are as we would expect and do not contain any major fragments. 4. optimization of the algorithm such a high number of points representing each particle would greatly affect the speed of the calculation. therefore, we focused on how we could reduce the figure 6. structure of elliptical voronoi particles before composition. figure 7. structure of elliptical voronoi particles after composition. number of points on the particle boundary while maintaining the accuracy of the boundary. therefore, we decided to see what effect replacing random points with specific points on the particle boundary. one of these characteristic points is the point of connection of the three cells. it is also the centre of the circle that touches the three cells. this problem is not trivial and is generally called the apollonius problem or the apollonius circle. f (x, y) = (x − xc)2 + (y − yc)2 = r2 (1) (xs − xci)2 + (ys − yci)2 = (rs − siri)2 (2) equation (1) is general expression of circle, where xc and yc refers to the center of the circle and r is radius. the following equation (2) represents the general solution of the apollonius circle, where si 113 jan vozáb, jan vorel acta polytechnica ctu proceedings figure 8. apollonius circles connecting circular particles. figure 9. optimized voronoi structure of circular particles before composition. is coefficient, that determines whether the resulting circle will be touched by the particle from the interior or the exterior. the result of this system of equations can be seen in the figure 8 for the same set of particles. the linkages between the particles are shown in blue. the red shows the circles between the particles as well as the points that these circles create in contact with particle the boundaries calculated by the unoptimized method are shown in black color. the figure 9 shows the voronoi diagram before composition. at first glance, one can see the much smaller number of cells of which it is composed. the figure 10 shows the resulting structure after composition on the resulting cells. small fragments can be seen where the cells are close together, or at the boundaries of the diagram when the particles have different radii. but the resultfigure 10. optimized voronoi structure of circular particles after composition. ing structure is formed at an average of 8 points per particle. even so, its accuracy is very consistent with the unoptimized method. we have tried a similar approach to calculate points for ellipses. but we encountered the problem here that the general equation of the apollonius circle is noticeably more complicated. with the general equation of the ellipse: f (x, y) = ((x − xc) · cos α + (y − yc) · sin α)2 a2 + ((x − xc) · sin α − (y − yc) · cos α)2 b2 = 1, (3) which then has to be converted into the general equation of the conic section: e(x, y) = ax2 + 2bxy + cy2 + 2dx + 2ey + f. (4) then, following the procedure defined in [10], several parameters are calculated to form a determinant, which for a single ellipse is defined as a polynomial of degree 184. the resulting system of equations: ∆1(xs, ys, s) = ∆2(xs, ys, s) = ∆3(xs, ys, s) = 0. (5) the resulting system of equations is a very challenging problem to solve. therefore, we tried to come up with a modification that would make the calculation possible. so we took inspiration from the circle equation, and considered that the ellipses will grow with one parameter until they intersect at one point. from this assumption comes the equation together with the system of equations: fi(x,y, r) = ((x − xci) · cos αi + (y − yci) · sin αi)2 (ai + r)2 + ((x − xci) · sin αi − (y − yci) · cos αi)2 (bi + r)2 − 1 = 0, (6) f1(x, y, r) = f2(x, y, r) = f3(x, y, r) = 0. (7) 114 vol. 40/2023 generation of ldpm structure formed by voronoi cells figure 11. apollonius circles connecting elliptical particles. the result of the system of equations is a polynomial of degree 27. it’s still not a simple solution, but it’s easy to solve the equations numerically. the resulting circles are shown in the figure 11 for the same set of particles. the figure 12 shows the structure and is made up of an average of 6 points per particle. the black line shows the boundaries for particles consisting of 44 points. the structure after composition is shown in red in the figure 13. there are clear fragments and differences from the exact boundary, but the optimized structure is very consistent even with a large saving of points 5. conclusion in this paper, the procedure of structure generation using the ldpm model is presented. it also describes how the structure can be generated using a voronoi diagram. this procedure relies on the representation of particles by points on their boundary for which the voronoi diagram is computed. this is followed by a composition of particles for which the structure already corresponds to their shape. then, a procedure is described on how the algorithm can be optimized using the specific representation. preliminary results have led us to the conclusion that the specifics are more telling in the construction of the resulting voronoi diagram and can be used to save a number of points on the particle boundary. this procedure will be more widely used in the future. it is mainly a matter of trying out additional specific points, then converting the structure to 3d. acknowledgements the financial support provided by the gačr grant no. 21-28525s and by the czech technical university in prague within sgs project with the application registered under the no. sgs20/155/ohk1/3t/11 is gratefully acknowledged. figure 12. optimized voronoi structure of elliptic particles before composition. figure 13. optimized voronoi structure of elliptic particles after composition. references [1] g. cusatis, a. mencarelli, d. pelessone, j. t. baylot. lattice discrete particle model (ldpm) for failure behavior of concrete. ii: calibration and validation. cement & concrete composites 33(9):891–905, 2011. mag id: 2094226000. https: //doi.org/10.1016/j.cemconcomp.2011.02.010 [2] f. aurenhammer. voronoi diagrams – a survey of a fundamental geometric data structure. acm computing surveys 23(3):345–405, 1991. mag id: 1967005434. https://doi.org/10.1145/116873.116880 [3] n. roussel. understanding the rheology of concrete. elsevier, 2011. [4] z. p. bažant, h. t. nguyen, a. abdullah dönmez. critical comparison of phase-field, peridynamics, and crack band model m7 in light of gap test and classical 115 https://doi.org/10.1016/j.cemconcomp.2011.02.010 https://doi.org/10.1016/j.cemconcomp.2011.02.010 https://doi.org/10.1145/116873.116880 jan vozáb, jan vorel acta polytechnica ctu proceedings fracture tests. journal of applied mechanics 89(6):061008, 2022. https://doi.org/10.1115/1.4054221 [5] b. l. karihaloo, a. carpinteri, m. elices. fracture mechanics of cement mortar and plain concrete. advanced cement based materials 1(2):92–105, 1993. https://doi.org/10.1016/1065-7355(93)90014-f [6] k. nagai, y. sato, t. ueda. mesoscopic simulation of failure of mortar and concrete by 3d rbsm. journal of advanced concrete technology 3(3):385–402, 2005. https://doi.org/10.3151/jact.3.385 [7] m. marcon, j. vorel, k. ninčević, r. wan-wendner. modeling adhesive anchors in a discrete element framework. materials 10(8):917, 2017. mag id: 2743109236. https://doi.org/10.3390/ma10080917 [8] q. du, v. faber, m. gunzburger. centroidal voronoi tessellations: applications and algorithms. siam review 41(4):637–676, 1999. mag id: 2051752778. https://doi.org/10.1137/s0036144599352836 [9] matlab. 9.11.0.1837725 (r2021b) update 2. the mathworks inc., 2021. [10] i. z. emiris, g. m. tzoumas. algebraic study of the apollonius circle of three ellipses. in eurocg, pp. 147–150. 2005. 116 https://doi.org/10.1115/1.4054221 https://doi.org/10.1016/1065-7355(93)90014-f https://doi.org/10.3151/jact.3.385 https://doi.org/10.3390/ma10080917 https://doi.org/10.1137/s0036144599352836 acta polytechnica ctu proceedings 40:111–116, 2023 1 introduction 2 ldpm structure 3 voronoi structure 4 optimization of the algorithm 5 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0050 acta polytechnica ctu proceedings 34:50–55, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague determination of water absorption coefficient of unfired earth materials different in used clay and ratio of components barbora mužíková∗, tereza plaček otcovská, pavel padevět czech technical university in prague, faculty of civil engineering, thakurova 7, 166 29 prague 6, czech republic ∗ corresponding author: barbora.muzikova@fsv.cvut.cz abstract. the paper is focused on unfired rammed earth and its water absorption properties. increasing the utilization of raw natural materials can be one of the approaches towards sustainable development. different prescriptions were designed and specimens were rammed. then they were put in the covering of nylon and settled in the box with soft foam that was moistened. the level of moistening was constant. the specimens were regularly weighted. specimens with montmorillonite clay have the highest values of the water absorption coefficient. montmorillonite clay has a higher binding capacity. the values are compared to values that were found in the literature. then the maximal capillary water capacity by area was determined. keywords: rammed earth, unfired earth, water absorption coefficient, clay. 1. introduction nowadays the price of building material is steeply rising and it begins to be lack of it. stone quarries have stocks for the next seven years in the czech republic and no new ones are opening. one way out of this problem is recycling and the second one is using natural local materials. this leads us to unfired earth that is local material, minimizing the cost for transportation. increasing the utilization of raw natural materials can be one of the approaches towards sustainable development. natural materials have a high potential of environmental quality for example in criteria: embodied co2 and so2 emissions, embodied energy, using renewable sources or easy recycling [1–3]. of course, using this material is not without any problems. since building with them was minimal in the last years, there are no sufficient data about its mechanical and physical data. these data are necessary for designing and building. figure 1. process of making rammed earth specimen. one of the problems for the unfired earth is water. the water inside or outside the earthen construction usually leads to problems. any time earth is to deal figure 2. schema of the water absorption test. with water needs to be controlled thoroughly: water content in mixture, vapor condensation in construction, or effects of weather. these are widely recognized issues of civil engineering. another big issue that is drawing unfired earth back from wide use is a strongly varying composition of raw material. the aim of this paper is to settle the water absorption coefficient for different clay mixtures and describe the influence of used clay and composition of the mixture on its water absorption characteristics [4–8]. 50 https://doi.org/10.14311/app.2022.34.0050 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 34/2022 water absorption coefficient of unfired earth figure 3. prescription of used specimens – graph shows percent procuration of clay, sand and water in the mixture. mixture sand/clay – w/c bulk density[%/%] – [-] [kg/m3] agl 6 85/15 – 0.400 2063 kr 3 80/20 – 0.400 2180 agl 1 80/20 – 0.370 2191 kr 1 80/20 – 0.370 2212 kr 8 80/20 – 0.290 2127 kr 7 80/20 – 0.250 2093 kr 11 75/25 – 0.400 2102 agl 2 75/25 – 0.370 2159 kr 2 75/25 – 0.370 2138 gem 3 75/25 – 0.295 1905 kr 13 70/30 – 0.370 2064 table 1. composition of mixtures and bulk density of specimens. 2. water in unfired earth water is very important for rammed earth it activates the binding properties of clay and thanks to it the material holds together. on the other hand, water is very dangerous to unfired earth. if loam becomes wet, it swells and changes from a solid to a plastic state [3]. mechanical properties are firmly affected by the moisture of the material. the unfired rammed earth is very sensitive to moisture and its structures can even collapse if the volume of water is high and affect the construction for a longer time period. 3. preparation of specimens the earth material is very sensitive to moisture and water that is why it is an important characteristic that needs to be examined. the structure of the earth is widely un-homogeneous, porous, and open. it is capable to absorb water and transport it in the material itself. rammed earth is a mixture of sand, clay, and water. clay works as a binder and water activate the binding forces of the clay. sand is a filler. the mechanical properties are depending on the composition of these components and their ratio and type of used clay. there are three basic kinds of clay – illite, kaolinite, and montmorillonite. eleven prescriptions were designed and made, they differ in the ratio of components and type of clay. prescriptions are shown in tab. 1. the prescriptions are arrayed by a rising amount of clay. the prescriptions are described as a ratio of sand and clay and the water ratio in the table t. water ratio is a relation between binder and water, in this case between clay and water. bulk densities for each prescription are shown in the table. the prescription are also shown in the fig. 3. the compounds are expressed by the percent. sand, clay, and water give together 100 %. in the tab. 1 just sand and clay give 100 % (and water is expressed by the water-clay ratio). for prescription agl the illite clay was used, for gem the montmorillonite clay and for kr mix of illite-kaolinite clay. the reciprocal ratio between the components is in percent. the orange color represents the clay, the yellow color represents sand and the blue color represents water. the first prescription agl 6 had the least amount of clay (14 %), the second group are specimens with clay containing around 19 % (kr 3, agl 1, kr 1, kr 8, kr 7). the third group is specimens with clay containing around 23 % (kr 11, agl 2, kr 2, 51 b. mužíková, t. p. otcovská, p. padevět acta polytechnica ctu proceedings figure 4. water absorption coefficient. figure 5. water absorption coefficient for earth mixture in literature [3]. gem 3) and the prescription kr 13 with the highest amount of clay 27 %. the production of specimens was as follows. firstly, the sand was mixed with two-thirds of water and mixed up. then the clay and the rest of the water were added. it was stomped by a drilling machine with a special ending. specimens of size 40×40×160 mm were prefabricated for the testing in the laboratory. the production was made by ramming into steel molds (shown fig. 1) by hand and by the drilling machine. the molds were wiped by oil and the earth was rammed in four to five layers. the last one was always made by hand. comprehensive strength and tensile strength in bending can be found in [9–11]. figure 6. specimens in nylon cover during testing. 4. method of determining water absorption coefficient specimens of measurement 40×40×160 mm were used for the test of water absorption. the absorbing area of specimens is 40×40 mm. specimens were covered in nylon to protect them from eroding. specimens were measured and weighted. a box with foam was filled with water until the foam was fully waterlogged. specimens were weighted up every hour, data were recorded and the condition of the specimen is judged. it means that a breakup of the specimen or magnitude of soaked up is observed. loam is a material with an open porous structure and it can transport water in capillaries. the water moves from the area with higher humidity to the area with lower humidity. capillarity means the capacity of water that can move such as was described. the 52 vol. 34/2022 water absorption coefficient of unfired earth figure 7. maximal capillary water capacity by area. figure 8. specimens after testing. process is called capillary action. the coefficient w is calculated by the formula: w = w √ t [ kg m2 · hod0.5 ] , (1) where w is the increase in weight per unit surface area and t is the time of measurement. w is calculated: w = m − md a [ kg m2 ] , (2) where m is the weight in the moment of measuring and md is the weight at the beginning of the test, the weight of the dry sample. the parameter w is also called the water movement and it represents the figure 9. type of clay – kaolinite, illite and montmorillonite. quantity of water that can be absorbed by a time period [3]. evaluated data of water absorption coefficient are shown in the fig. 4. the averaged value and its standard deviation are shown. 53 b. mužíková, t. p. otcovská, p. padevět acta polytechnica ctu proceedings 5. method of determination maximal capillary water capacity area the specimens were settled in the water bath as was described above. they were in the water until they broke up. the maximal weight was recorded and then the maximal capillary water capacity was determined by the formula: wmax = mmax − md a [ kg m2 ] , (3) where mmax is the maximal weight until the specimen was broken. the maximal weights are not measured at the same time, the time differed for prescriptions. 6. evaluation of the results 6.1. water absorption coefficient calculated values of water absorption coefficients are shown in the fig. 4. the prescriptions are arrayed by a rising amount of clay. water absorption coefficient w says how quickly can the material absorb the water. the evaluation is after one hour of testing and the absorbing area of specimens was 40×40 mm. the highest value has the prescription gem 3 with montmorillonite clay. this type of clay has the highest absorption capabilities. the value is 21±7 kg/m2hod0.5. similar value have solid brick 25.5 kg/m2hod0.5 according to [3]. the minimum value had prescriptions with a mixture of illitekaolinite and illite clay. prescription agl 6, kr 8 and kr 7 have higher values than others, they have less water just 5 % and 6 %. other prescriptions had similar value around 3 kg/m2hod0.5. the value responds to the values that were found by g. minke in [3] for his earth mixture. there were also values around 3 kg/m2hod0.5. these values were for different types of loams. these values are shown the fig. 5 with bulk densities of specimens. they can be compared with measured values in the tab. 4. the main influence has the type of used clay, the amount is not so important as the type. clay minerals are three basic ones – kaolinite, illite and montmorillonite. they differ in structure. clays have hexagonal lamellar crystalline structures. these lamellas consist of different layers that are usually formed around silicon or aluminum cores [3]. this is shown in the fig. 9. the difference between clays can be seen in this picture. kaolinite is a two-layered mineral and has a lower ion-binding capacity. on the other hand, montmorillonite is a three-layered mineral and has a higher ion binding capacity [4]. 6.2. maximal capillary water capacity by area calculated values of maximal capillary water capacities by area are shown in the fig. 7. the prescriptions are arrayed by a rising amount of clay again. the maximal capillary water capacity by area is determined from the maximal weight at the end of testing before the specimens collapsed. the time is not the same at all prescriptions, some of them collapsed sooner than others. the results were divided in three groups the first one with values around 30–35 kg/m2 (agl 1, kr 1, kr 8, agl 2, kr 2, gem 3). the second group are prescriptions with value 20–25 kg/m2 (agl 6, kr 3, kr 7 and kr 11) and the minimal value had kr 13 108 kg/m2. gem 3 with montmorillonite clay was specifed. the volume of the specimen extended a lot in the wet bottom and then it brooked. no other mixture had expansion like the mixture with montmorillonite clay. 7. conclusions the price of building material is steeply rising and it begins to be lack of it. one way out of this problem is recycling and the second one is using natural local materials. this leads us to unfired earth that is local material, minimizing the cost for transportation. increasing the utilization of raw natural materials can be one of the approaches towards sustainable development. the earth material is very sensitive to moisture and water that is why it is an important characteristic that needs to be examined. the structure of the earth is widely un-homogeneous, porous, and open. it is capable to absorb water and transport it in the material itself. it is important to search the properties connected with water. different prescriptions were designed and specimens were rammed. then they were put in the covering of nylon and settled in the box with soft foam that was moistened. the level of moistening was constant. the specimens were regularly weighted. the highest values of water absorption coefficient have the specimens with montmorillonite clay that has the higher binding capacity. the values are compared to values found in the literature. then the maximal capillary water capacity by are was determined. the results of maximal capillary water capacity by area were divided into three groups the first one with values around 30–35 kg/m2 (agl 1, kr 1, kr 8, agl 2, kr 2, gem 3). second group are prescriptions with value 20–25 kg/m2 (agl 6, kr 3, kr 7 and kr 11) and the minimal value had kr 13 18 kg/m2. list of symbols a area [ m2] m weight of specimen in the moment of measuring [kg] md weight of specimen in the begging of the test (dry sample) [kg] mmax maximal weight of specimen [kg] t time of testing [hod] w water absorption coefficient [kg/m2hod0 5] 54 vol. 34/2022 water absorption coefficient of unfired earth w capillary water capacity [kg/m2] wmax maximal capillary water capacity [kg/m2] acknowledgements the research was financially supported by the czech science foundation (gacr no.18-0884s) and by the faculty of civil engineering at ctu in prague (sgs project no.sgs19/148/ohk1/3t/11). references [1] f. havlik. development and experimental verification of mechanical-physical properties of pre-formed rammed earth wall panel, vol. 1. cvut, prague, 2017. [2] j. norton. building with earth, a handbook, vol. 1. intermediate technology development group limited, london, 1986. [3] g. minke. building with earth design and technology of sustainable architecture, vol. 2. de gruyter, berlin, 2009. [4] i. zabickova. hlinené stavby. era 21, brno, 2002. [5] m. rauch. refined earth construction and design with rammed earth, vol. 1. detail, munich, 2015. [6] kianfar, ehsan, v. toufigh. reliability analysis of rammed earth structures. construction and building materials (127):884–895, 2016. [7] h. araki, j. koseki, t. sato. tensile strength of compacted rammed earth materials. soils and foundation (56(2)):189–204, 2016. [8] l. miccoli, d. v. oliveira, r. a. silva, u. muller. static behaviour of rammed earth: experimental testing and finite element modelling. materials and structures (48(10)):43–56, 2015. [9] t. otcovska, p. padevet. dependence of tensile bending strength of rammed earth on used clay composition and amount of mixture water. modern methods of experimental and computational investigations in area of construction ii pp. 48–53, 2017. [10] t. otcovska, p. padevet. dependence of compressive strength of rammed earth on used clay composition. experimental stresss analysis 2016 pp. 107–108, 2016. [11] i. zabickova, t. otcovska, p. padevet. compressive strength of unburned clay masonry. modern methods of experimental and computational investigations in area of constructions pp. 31–34, 2016. 55 acta polytechnica ctu proceedings 34:50–55, 2022 1 introduction 2 water in unfired earth 3 preparation of specimens 4 method of determining water absorption coefficient 5 method of determination maximal capillary water capacity area 6 evaluation of the results 6.1 water absorption coefficient 6.2 maximal capillary water capacity by area 7 conclusions list of symbols acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0022 acta polytechnica ctu proceedings 40:22–26, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague treatment of polypropylene microfibers by atmospheric and low-pressure plasma – application to a reinforced cement composite containing recycled concrete jakub ďurejea, ∗, zdeněk prošeka, jan trejbala, štěpán potockýb, radim hlůžeka a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b czech academy of sciences, institute of physics, cukrovarnická 10, 162 53 prague 6, czech republic ∗ corresponding author: jakub.dureje@fsv.cvut.cz abstract. the effect of atmospheric and low-pressure plasma modification on polypropylene (pp) microfibers was examined. mechanical changes on the microfiber surfaces were observed using scanning electron microscopy (sem). next, wettability was measured using the packed-cell method. the fibers were applied into a cement matrix containing micro-milled recycled concrete. test specimens were made and then the dynamic modulus of elasticity was continuously measured. after 28 days were made in the test specimens central notches to a depth of 14 mm. finally, bending tests were performed. from the results, the fracture energy of the composite material was calculated. it was proven that low-pressure plasma modification as well as atmospheric plasma modification increases the wettability of pp fibers with water. furthermore, it was found that samples containing plasma-modified microfibers have a higher fracture energy compared to the same samples with fibers without plasma modification. conversely, plasma modification had no effect on the dynamic modulus of elasticity. keywords: atmospheric plasma, plasma modification, plasma treatment, polypropylene microfiber, oxygen plasma, wettability, fracture energy, sem. 1. introduction nowadays, proper waste management is very important. old waste concrete can be recycled and subsequently reused. it is most often used as a bottom layer for roads. in this work, recycled concrete was ground using a high-speed mill and then used as a filler for the cement composite material. the advantage of this use is possibility to use all fractions of recycled concrete and possibility to activate some non-hydrated cement in recycled concrete [1, 2]. the cementitious composite material was reinforced using polypropylene (pp) microfiber reinforcement. the properties of the resulting composite material depend, among other things, on the interfacial transition zone (itz). in this zone occurs interaction between the cement matrix and the fibers [3]. to improve the adhesion between the fiber and the matrix, it is possible to modify the surface of the fibers, which will lead to an improvement of the mechanical properties final composite material. using plasma, we can modify the fiber surface both mechanically and chemically [4]. the mechanical effect of plasma modification is caused by ion bombardment. the chemical effect of plasma modification is mainly caused by the many chemical groups that are generated during modification [5, 6]. the effect of plasma modification on microfibers surface depends on many parameters, including time, gas, power, and device. plasma treatments are performed in a low-pressure chamber or at atmospheric pressure [7, 8]. 2. materials and specimens portland cement, micro-milled recycled concrete, plasma treated polypropylene (pp) microfibers and water were used to produce the test samples. the ratio of cement to recycled was 1:1. the water ratio w/c+r was 0.32 for each sample. portland cement cem i 42.5 r radotín (českomoravský cement, czechia) was used. micro-milled recycled concrete (lavaris, czechia) was made from concrete drainage gutters using a high-speed mill. the specific surface of micro-milled recycled concrete is around 36 m2/kg. microfibers fibrofor multi (contec fiber, switzerland) were made of polypropylene. the fibers are made in bundles (type 127), the diameter of individual filament is about 32 µm and the microfiber length is 12 mm ±5 %. the surface of the microfibers was modified using atmospheric or low-pressure plasma. low-pressure plasma treatment was performed by tesla vt214 device using an rf source – gas pressure in the chamber of device was 20 pa. atmospheric pressure modification was performed by roplass rps 400 device using a dielectric barrier discharge. the process parameters of oxygen low-pressure plasma modification were chosen based on previous experiments [9]. 22 https://doi.org/10.14311/app.2023.40.0022 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 40/2023 treatment of polypropylene microfibers . . . set device input power [w] time [s] gas pressure [pa] ref a roplass rps 400 300 480 (4×120) air atmospheric t tesla vt213 100 480 (2x×240) oxygen 20 table 1. process parametrs of plasma modification. set cement [g] recyclate [g] water [g] w/c+r microfibers [g] microfibers [%] ref 1500 1500 960 0.32 28.0 2 a 1500 1500 960 0.32 28.0 2 t 1500 1500 960 0.32 28.0 2 table 2. composition of the samples. the oxygen low-pressure plasma modification of the microfibers was performed for a total of 480 s, while the process was paused after 240 s, the fibers were mixed and then the process was started again. for the same duration, the fibers were modified in atmospheric pressure plasma, where the working was atmospheric air. to achieve a uniform modification of the surface of the fibers, the fibers were mixed during the process after every 120 s. the process parameters of plasma modification are in table 1. a total of three sets specimens were made, each set containing six test specimens. the dimensions of the test specimens were 40 × 40 × 160 mm. samples were unmolded 24 hours after production and were stored 28 days in standard laboratory environment. composition of the samples is shown in the table 2. 3. experimental methods fiber surface was examined by scanning electron microscopy (sem). first, a thin layer of platinum was sputted on the surface of the fibers using a mini sputter coater quorum sc7620 at a pressure of 8 pa. subsequently, the surface of the fibers was examined using a merlin zeiss sem. the surface of the fibers was examined at a magnification of 20,000× (figure 1). the wettability of the fibers was measured by the packed-cell method. the fibers were insert into a container with a perforated bottom, after that this container was immersed into water for 60 seconds. container with the fibers was weighed on the laboratory scale before immersion and 120 seconds after immersion. finally, the percentage weight of water to weight of fibers was calculated (1) [10]: mv = (mm − mn) − (ms − mn) (ms − mn) · 100, (1) where mv the weight of water to the weight of fibers ratio [%], mm the weight of wet fibers and packed-cell measuring set [g], (a). plasma modification in low-pressure by oxygen. (b). plasma modification in atmospheric pressure by air. figure 1. sem image – microfiber surface after modicifation. ms the weight of dry fibers and packed-cellmeasuring set [g], mn the weight of packed-cell measuring set [g]. the dynamic modulus of elasticity of the samples was continuously measured by the resonance method (brüel&kjær, denmark). the modulus of elasticity was measured 7, 14, 21 and 28 days after samples production. the measuring system includes an impulse hammer brüel&kjær type 8206, response sensor brüel&kjær type 4519-003 and measure device brüel&kjær front-end 3560b-120. the dynamic mod23 j. ďureje, z. prošek, j. trejbal et al. acta polytechnica ctu proceedings figure 2. position of the sensor and the impulse hammer on the sample to measure the fundamental frequencies from longitudinal (left), transverse (center), and torsional (right) oscillations; s – sensor, b – impact hammer [11]. ulus of elasticity was determined from longitudinal, transverse and torsional oscillations (figure 2). finally, the dynamic modulus of elasticity was calculated using pulse labshop software version 14.0.1. for measurement of fracture energy was performed a notch in the middle of the length test specimens. notch was performed using an automatic saw with a water-cooled diamond blade (achilli, italy). the depth of the notch was 14 mm, which is approximately one third of the height of the specimen. the notch width was 3 mm. subsequently, a three-point bending test was performed. samples were loaded by electromechanical press (mts, usa). samples were loaded with constant displacement at a speed of 1.5 mm/min. the fracture energy was calculated using the formula (2): gf = af bw , (2) where gf fracture energy [j/m2], af the work of loading force [j], bw the area of the crack ligament [m2]. the work of the loading force was calculated as the integral of the function from the force-displacement graph: af = ∫ smax 0 f ds, (3) where s displacement during loading test, f force during loading test. 4. results and discussion in the sem images were observed on the fiber surfaces significant mechanical changes compared to the reference fibers for both types of modification. however, the changes on the fiber surfaces are different for each type of modification. fibers modified by low-pressure oxygen plasma have holes on their surfaces. fibers modified by atmospheric pressure plasma have pimples (drops) on their surfaces. the wettability of the figure 3. weight of water to weight of microfibers. figure 4. dynamic modulus of elasticity from 0 to 28 days. fibers increased after both modifications, the amount of water between the fibers increased in both cases by approximately 20 % compared to the reference values. the chemical effect of plasma modification is approximately the same in both cases (figure 3). the dynamic modulus of elasticity was slightly lower for the samples containing plasma-modified fibers than the reference samples. for fibers modified by atmospheric plasma, modulus of elasticity decreased by approximately 1 %, for fibers modified by low-pressure oxygen plasma decreased by 3.5 % (figure 4). decrease in modulus of elasticity is negligible, most likely it was 24 vol. 40/2023 treatment of polypropylene microfibers . . . figure 5. flexural test – force-displacement graph. caused by slightly worse workability of the cement mixture. that corresponds to a slightly lower density of samples with plasma-modified fibers compared to the reference samples. the fracture energy was higher for the samples with plasma-modified fibers compared to the reference samples. for samples containing fibers modified by atmospheric pressure plasma, it was an increase of approximately 8 %, for samples containing fibers modified by low-pressure oxygen plasma, it was an increase of approximately 25 % (figures 5, 6). 5. conclusion the surface of the fibers was modified both mechanically and chemically. based on the experiment, we can conclude that: • modification by oxygen low pressure plasma caused mechanical changes on the fibers surfaces. there were observed by sem (magnification 20,000×) holes caused by this modification. • modification by atmospheric pressure plasma caused mechanical changes on the fibers surfaces. on the fiber surface were observed by sem (magnification 20,000×) formations that look like drops or pimples caused by this modification. • the wettability increased approximately the same after both modifications. the amount of water between the fibers increased by approximately 20 % in both cases. • modifications had almost no effect on the modulus of elasticity. the modulus of elasticity slightly decreased, for samples containing microfibers modified by atmospheric plasma decreased by 1 %, for samples containing microfibers modified by lowpressure oxygen plasma decreased by 3.5 %. this decrease was probably caused by the slightly worse workability of the fresh cement composite mixture for the samples containing plasma-modified fibers compared to the reference samples. • the fracture energy of samples with plasmamodified fibers increased in both cases of modification. for samples modified by atmospheric pressure plasma, it increased by 8 %, for samples modified by low-pressure oxygen plasma, it increased by 25 %. figure 6. fracture energy of samples. the increase in fracture energy was mainly caused by the mechanical effect of plasma modification. the tested plasma modifications succeeded in increasing the fracture energy of the cement composite material containing micro-milled recycled concrete. both tested modifications were suitable. samples containing fibers modified by low-pressure oxygen plasma had a higher fracture energy than samples containing fibers modified by atmospheric pressure plasma. on the other hand, plasma modification performed in atmospheric pressure is significantly easier to apply in mass production compared to modification performed in low-pressure. the workflow and devices for lowpressure plasma are more complicated compared to atmospheric plasma devices. acknowledgements this work was financially supported by the czech technical university in prague – the project sgs22/089/ohk1/2t/11. the authors also thank the center for nanotechnology in civil engineering at the faculty of civil engineering, czech technical university in prague. references [1] k. mcneil, t. h. k. kang. recycled concrete aggregates: a review. international journal of concrete structures and materials 7(1):61–69, 2013. https://doi.org/10.1007/s40069-013-0032-5 [2] j. topič, z. prošek. properties and microstructure of cement paste including recycled concrete powder. acta polytechnica 57(1):49–57, 2017. https://doi.org/10.14311/ap.2017.57.0049 [3] g. prokopski, j. halbiniak. interfacial transition zone in cementitious materials. cement and concrete research 30(4):579–583, 2000. https://doi.org/10.1016/s0008-8846(00)00210-6 [4] f. d. egitto, l. j. matienzo. plasma modification of polymer surfaces for adhesion improvement. ibm journal of research and development 38(4):423–439, 1994. https://doi.org/10.1147/rd.384.0423 [5] f. awaja, s. zhang, n. james, d. r. mckenzie. free radicals generated by ion bombardment of a semi-crystalline peek surface. plasma processes and 25 https://doi.org/10.1007/s40069-013-0032-5 https://doi.org/10.14311/ap.2017.57.0049 https://doi.org/10.1016/s0008-8846(00)00210-6 https://doi.org/10.1147/rd.384.0423 j. ďureje, z. prošek, j. trejbal et al. acta polytechnica ctu proceedings polymers 9(2):174–179, 2012. https://doi.org/10.1002/ppap.201100056 [6] e. couto, i. h. tan, n. demarquette, et al. oxygen plasma treatment of sisal fibers and polypropylene: effects on mechanical properties of composites. polymer engineering & science 42(4):790–797, 2002. https://doi.org/10.1002/pen.10991 [7] c. mandolfino. polypropylene surface modification by low pressure plasma to increase adhesive bonding: effect of process parameters. surface and coatings technology 366:331–337, 2019. https://doi.org/10.1016/j.surfcoat.2019.03.047 [8] m. shenton, m. lovell-hoare, g. stevens. adhesion enhancement of polymer surfaces by atmospheric plasma treatment. journal of physics d: applied physics 34(18):2754, 2001. https://doi.org/10.1088/0022-3727/34/18/307 [9] j. ďureje, z. prošek, j. trejbal, et al. influence of oxygen and argon plasma treatment on wettability and surface morphology of polypropylene microfibers. acta polytechnica ctu proceedings 34:11–14, 2022. https://doi.org/10.14311/app.2022.34.0011 [10] j. ďureje. plasma modification of microfiber reinforcement to improve adhesion with cement matrix containing recycled concrete. master’s thesis, czech technical university in prague, prague, 2018. [11] čsn 73 1372. nedestruktivní zkoušení betonu – rezonanční metoda zkoušení betonu, 2012. 26 https://doi.org/10.1002/ppap.201100056 https://doi.org/10.1002/pen.10991 https://doi.org/10.1016/j.surfcoat.2019.03.047 https://doi.org/10.1088/0022-3727/34/18/307 https://doi.org/10.14311/app.2022.34.0011 acta polytechnica ctu proceedings 40:22–26, 2023 1 introduction 2 materials and specimens 3 experimental methods 4 results and discussion 5 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0630 acta polytechnica ctu proceedings 38:630–634, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague rescue of a stone bridge with respect to current condition and standards petr řeřichaa, ∗, petr fajmana, lucy davisb a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 22, 166 29 praha, czech republic b mcgill university, faculty of engineering, 45 sherbrooke street west, montreal, canada ∗ corresponding author: petr.rericha@fsv.cvut.cz abstract. a three span stone masonry bridge dating back to the middle of the 19th century, still in the roadway network, is assessed to prevent its demolition. its industrial heritage value and ecological concerns were the principal reasons for the assessment. the carbon footprint of the stone arch replacement would be approximately 200 t co2 emission owing to 430 m3 reinforced concrete in the new structure. besides the cultural monument, considerable energy, co2 emissions and natural resources can be saved in accordance with the sustainable development goals. standards, guides and commercial software do not provide an adequate support for the assessment of masonry arch bridges, therefore an innovative two phase application of a commercial linear sttructural analysis code and an original in-house code was developed for the purpose. thousands of stone ach bridges still in service worlwide give the case study more general importance. keywords: natural resources preservation, sandstone, bridge, arch, stress, displacement. 1. introduction the three-span sandstone arch bridge was built in 1857 and remains an important part of the second-class roadway network, see figure 1. it was considered for a reconstruction with total demolition of the arches. the owner planned to remove the arches and top parts of the piers and replace them with a reinforced concrete structure that would be a reminder of the arches of the stone bridge. in order to prevent an irrecoverable loss of the industrial heritage, an extensive load capacity assessment was conducted for the bridge with the replacement of the pavement. standard assessments determine load rating based on the limit states analysis. the load models used for the purpose differ in various countries substantially, for instance either the ultimate (in europe) or serviceability limit states (in the us) are preferred. instead, the check of the available building standard’s criteria is used herein. the conditions of the en 1990 part 2 traffic loads on bridges have been checked to testify that the bridge with its valuable stone arches can remain a part of the roadway network without limitations. a number of clauses of the en 1996-1-1 could be applied even though the standard is not designed for bridges. “there is a multitude of assessment methods available for masonry arch bridges, yet there is no widely accepted framework for their application”, this quotation from [1] is a sad fact. tens of papers report on sophisticated methods based on advanced material and structural models. none of them are acceptable for the design practice for two principal reasons, figure 1. front view of the sandstone bridge at ponikla. (a) the application requires specific high competencies of the designer and specialized software (b) the input material and structure data are not available for the models. the present approach is simple: the assessment method can use only the data available in a common design environment and with reasonable cost. this admits relatively simple material models. the finite element technology is preferred since it can model the components of the bridge and their interactions in a way that other methods and codes (modified mexe, plasticity theorems like ring [2], castiligliano theorem like ctap code [3] and others) cannot. at the same time it is widely available. the absence of design standards or other “widely accepted framework” gives the load rating of an important bridge the character 630 https://doi.org/10.14311/app.2022.38.0630 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 rescue of a stone bridge with respect to current condition and standards figure 2. new design of the bridge. of a research task. an assessment method was developed in the course of the solution of the bridge. the objective of the contribution is to outline the principal steps of the method and its application to similar bridges. the method is tailored to masonry bridges management practice and demostrates its application. the sample solution documents the potential of the method for sustainable infrastructure management and ecological impact. 2. the original restoration concept, ecological impact scheduled for reconstruction, the owner planned to remove the arches and top parts of the piers and replace them with a reinforced concrete structure that would remind the arches of the stone bridge, see figure 2. however, considerable economical savings and benefits for the environment can be achieved when the arches are preserved. the construction of the new reinforced concrete structure would require the destruction and deposition of 230 m3 high quality stone masonry built with craftsmanship which would be lost forever. removal and deposition of 400 m3 backfill would cost money and ecological load. the carbon footprint would be left owing to 430 m3 reinforced concrete with approximately 200 t of co2 emission. the most important offense against sustainable building would probably be the life span of the new reinforced concrete structure. the stone arches are 170 years old and still in service in spite of the neglected maintenance in the last decades. with adequate protection against water leaks they can serve for 150 more years. the service life of the proposed reinforced concrete structure willl certainly be substantially shorter, as in particular the struts above the piers with notch hinges are susceptible to early deterioration. the ecological aspects of the reconstruction speak for the preservation of the stone arches. 3. load capacity of stone arch bridges 3.1. material models the dominant material failures and limit states in the stone arch bridges are (a) the tension cracking in the bed joints, (b) the compression failure in the bed joints. the former occurs already at the service load levels and may result in gradual development of virtual hinges and stability loss of the arch with the growing live load. it must be represented in the material model in order to assess the redistribution of the internal forces. the trouble with this limit state is that there are no generally accepted quantitative criteria worldwide. recently the czech guide tp199 of the ministry of transport of the czech republic introduced an admissible relative depth of the cracks in the bed joints. the other trouble is that the predestined orientation of the cracks makes the no-tension models in general purpose codes almost unusable. this failure mode can effectively be included only in beam elements of the 2d models or shell elements of the 3d models where the kinematic assumptions and dominant cross-section normal stresses facilitate the specific no-tension model. the thick shell and beam finite elements with the no-tension property are derived in [4]. the latter limit state is a kind of a brittle failure, partially contained in case of the arch bridge with fill. the failure criterion for this limit state is the compression strength of the masonry. this property is the only one that can be determined in a diagnostic survey with acceptable cost and reliability. en 1996-1-1 provides support for determining the masonry compression strength from the compression tests of the core bored specimens of the stones and by nondestructive boring tests of the mortar (pzz 01 device). the tests were performed for the ponikla bridge masonry arches and the compression design strength was determined fd = 2 mpa. the value is rather conservative since the sandstone ashlars of the vault are in good condition. the backfill and pavement are linear elastic. in order to eliminate extensive backfill failures, the drucker/prager criterion is checked in the fill elements and indicated by asterisks in the graphic output. the elastic properties used in the fem models are summarized in table 1. they are deliberately conservative for the backfill. a review of the backfill properties was conducted in the published literature [5] and these are the values at the low end. during the method development, a parametric study showed that softer fill always implies worse stresses in the vault drawing 631 p. řeřicha, p. fajman, l. davis acta polytechnica ctu proceedings figure 3. the minimum stresses [kpa] and crack depths in the ponikla central span arch. the displacements are 1000 times scaled up. e [gpa] g [gpa] ν ρ [t/m3] arch 10 5 0.25 2.0 backfill 0.1 0.043 0.15 1.8 pavement 31 12.917 0.2 2.5 spandrel 0.1 0.04 0.25 2.2 piers/abutments 10 4 0.25 1.8 table 1. elastic properties considered in this case study. an important conclusion. the fill material constants from the table are on the safe side and can be used in bridge assessment when nothing is known about the fill quality. the value 0.4 gpa is recommended when at least a compacted cohesionless material can be assumed. selection of the young moduli of the arch parts is facilitated by a convenient property of the fem model, see the next section [6]. the short summary of the material models and properties exposes the uncertainties involved which inevitably imply low reliability of the analytical models of the masonry arch bridges. on the other hand, most of them were built in the 19th century according to empiric rules with ample reserve of load capacity. decades of their service testify to their robustness and reliability. these aspects compensate the low reliability of the analytical models, asserts boothby in [7]. 3.2. finite element bridge models it is difficult to account in a single complex model for the interaction of the spans and piers, the transverse asymmetry of the load and structure and non-linear arch behavior. the assessment method therefore uses different fem models for the tasks. the checks required by the standards are performed in a non-linear 2d (plane strain condition) model of a single span featuring the no-tension material specified above. the simplest timoshenko beam elements [4] are used for the stone arch so that the normal stresses in the cross-section planes are used in the above failure conditions. the backfill is meshed by standard cst triangles and the pavement by linear timoshenko beam elements. the meshing is shown in figure 3. clamped arch springings are assumed. the results are slightly non-conservative owing to the plane strain 2d idealization which neglects all transverse effects. correction factors are used to make up for the error, see the next paragraph. the linear elastic no-tension material model secures robust convergence of the equilibrium iterations so that unskilled users can use the model and code. the model also implies that stresses and crack depths do not change when the young moduli of all materials are multiplied by a common factor. a non-commercial code was developed for the purpose and will be available as an octave/matlab script when a minimum user interface is completed. the other two models are a 3d linear model of one span with the same boundary conditions as in the non-linear model and a 3d model of the whole bridge including piers and abutments. they are primarily used to assess the differences between the 2d and 3d models and the interaction of the arches, piers and abutments.the differences between the 2d and 3d models are expressed in terms of correction factors for the 2d solutions. the factors are then used to correct the results of the non-linear model. this is not quite consistent but the bed joints cracking is localized to relatively small volumes of the vault and the imperfection is thus accepted. a front view of the bridge and its 3d fem model are shown in figures 1 and 4. the spans are all approximately 11.4 m and the full width of the bridge is 7.6 m. the single span 3d model is an extraction of the full bridge separated by adjacent pier’s symmetry planes. on the planes, the symmetry boundary conditions are applied (zero normal and free in-plane 632 vol. 38/2022 rescue of a stone bridge with respect to current condition and standards figure 4. full 3d bridge model with two tandem axle forces. displacements). the elements are based on isoparametric bricks/quadrilaterals enriched by rotational degrees of freedom at all nodes. the commercial package rfem from dlubal software [8] was used in the sample solution. any other package can be used with an adequate modeler. the model’s build itself is rather easy since details of the sructure can be omitted. extraction and processing of the relevant stresses and other data may be more demanding and depends largely on the available software. the other purpose of the full 3d model is to asses the interaction of the spans, piers and abutments. the dominant effect of the interaction is the widening of the arch subject to the local live load, see section 3.3 for the live load model. the vertical displacements and rotations of the arch springings owing to the interaction are neglected in the method. the arch widening is obtained as the difference of the averaged horizontal displacements of the right and left springings. it is then used in the non-linear 2d model as an imposed displacement (kinematic load). 3.3. load model the load model lm1 of the en 1991-1-1 is used. it is known to be critical for bridges with these space parameters. the same position of the two axle tandems is considered in all calculations, as shown in figure 1. the position above the quarter span is generally acknowledged to be the most adverse for the arch. the uniform continuous live loads of lm1 are not shown in figure 4 but applied on the loaded half of the span in the calculations. it is worth mentioning that lm4 class 1800 kn of the en standard gives approximately the same total load on the half span but is almost uniformly distributed. the dead load is included via mass densities of the bridge components materials. the load factors of the ultimate limit state are used according to the en and national application document. loading in the 2d model is taken as the average over the bridge width of the 3d model loads. 3.4. results for the design load of the ponikla central span the solution of the ponikla bridge for the design loads of the en 1991-2 is shown in figure 3 with axle tandems of the lm1 allocated at a quarter span of the arch. the dead load of the arch, fill and pavement is accounted for. the third load component is the imposed horizontal displacement 0.13 mm of the right arch springing, fill boundary and pavement end. it equals the arch widening obtained in the full 3d linear model, see the last paragraph of section 3.2. the absolute values of the minimum stresses in the arch elements are inserted at the arch face where they occur (the maximum stresses are always nonnegative). the same quantities are inserted at the lower and upper faces of the pavement although they are irrelevant with respect to the arch assessment. the 0.2 m thick reinforced concrete pavement is considered linear elastic in the model. the bending “waves” characteristic for beams on an elastic layer can be observed in the figure, note the alternating faces of the minimum stresses in the pavement elements. the crack depths in the arch elements are indicated by the inserted line segments. the crack mouth widths in the bed joints can be computed from the crack depths, the voussoir lengths and the minimum strains in the elements when the timoshenko kinematic assumption is employed. the arch model is a 1d continuous beam, where the arch elements lengths depend on the chosen mesh density. the voussoirs are approximately 0.4 m long whereas the elements just about 0.33 m. the load intensities and material properties used in the non-linear plane strain 2d model are averages 633 p. řeřicha, p. fajman, l. davis acta polytechnica ctu proceedings cross-section locations 0 0.25 0.5 extrados stress correction factor 1.18 1.08 1.1 intrados stress correction factor 1.05 0.94 table 2. the 3d/2d correction factors for the compressive stresses at three cross-section locations. across the vault width 7.6 m. however, the load model lm1 for two lanes is not symmetric with respect to the longitudinal symmetry plane of the bridge. the asymmetry effect is evaluated through comparison of the 3d model solutions for the true lm1 load distribution with the solution for the average uniform distribution in the transverse direction. the ratios of the two stress’ sets are used as 3d/2d correction factors and are summarized in table 2 for the four cross-section locations along the arch span. only compressive stresses are considered. the stress values from figure 3 are multiplied by the correction factors from table 2 in the last step of the method. the assessment of the bridge using the method determined its safety in the ultimate limit state as defined in en 1991-1. 4. conclusion a method for the assessment and load rating of the stone masonry arch bridges is presented. it is designed for application in bridge management. a single material property, the vault masonry compression strength, is necessary. users need no special skills beyond adequate experience with the application of the fem codes. the safety assessment of the ponikla bridge using available standards together with activities of other institutions and individuals in the cultural heritage protection helped to avert the demolition of the bridge’s arches. the simple material model and the dedicated finite element can be used to check the complex structure against available standards conditions. an extension of the concept to 3d shells is currently under way and then the parallel linear 3d model will not be necessary for the application of the method. preservation of the ponikla bridge represents a contribution to the responsible consumption and production goal of the sustainable development policy. acknowledgements the authors acknowledge the support of the project naki dg20p02ovv001 (tools for the preservation of historical values and functions of arch and vaulted bridges) provided by the ministry of culture of the czech republic. references [1] n. gibbons, p. j. fanning. ten stone masonry arch bridges and five different assessment approaches. in arch’10 6th international conference on arch bridges, pp. 482–489. fuzhou, 2010. [2] m. gilbert. ring: a 2d rigid block analysis program for masonry arch bridges. in proceedings of the arch01: 3rd international arch bridges conference, pp. 459–464. paris, 2001. [3] r. j. bridle, t. g. hughes. ctap theory. journal inst highways transp 36:21–22, 1989. [4] p. řeřicha. a mindlin shell finite element for stone masonry bridges with backfill. acta polytechnica 62(2):293–302, 2022. https://doi.org/10.14311/ap.2022.62.0293 [5] l. j. davis. load rating of a masonry arch bridge: comparison of 2d and 3d finite element models. master’s thesis, czech technical university in prague, faculty of civil engineering, 2021. [6] p.-o. persson, g. strang. a simple mesh generator in matlab. siam review 46(2):329–345, 2004. https://doi.org/10.1137/s0036144503429121 [7] t. e. boothby. empirical design of masonry arch bridges. journal of architectural engineering 26(1):02519002, 2020. https: //doi.org/10.1061/(asce)ae.1943-5568.0000388 [8] rfem 6, fem structural analysis software. [2021-0416]. https://www.dlubal.com/en/products/rfem-feasoftware/what-is-rfem 634 https://doi.org/10.14311/ap.2022.62.0293 https://doi.org/10.1137/s0036144503429121 https://doi.org/10.1061/(asce)ae.1943-5568.0000388 https://doi.org/10.1061/(asce)ae.1943-5568.0000388 https://www.dlubal.com/en/products/rfem-fea-software/what-is-rfem https://www.dlubal.com/en/products/rfem-fea-software/what-is-rfem acta polytechnica ctu proceedings 38:630–634, 2022 1 introduction 2 the original restoration concept, ecological impact 3 load capacity of stone arch bridges 3.1 material models 3.2 finite element bridge models 3.3 load model 3.4 results for the design load of the ponikla central span 4 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.35.0001 acta polytechnica ctu proceedings 35:1–7, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague kalker’s coefficient c11 and its influence on the damping and the retuning of a mechanical drive torsion system of a railway vehicle vojtěch dybala czech technical university in prague, faculty of mechanical engineering, department of automotive, combustion engines and railway engineering, technická 4, 160 00 prague 6, czech republic correspondence: vojtech.dybala@fs.cvut.cz abstract. within the research of electromagnetically excited torsion oscillations in the mechanical part of traction drive systems of modern railway vehicles, which has been realized at the faculty of mechanical engineering at the ctu in prague, there are two separate simulation models in use. the basic calculation model, which is utilized to gain basic characteristics of the torsion system as natural frequencies and natural modes of oscillations. and the complex simulation model, which simulates a drive of the vehicle. this contribution is focused on the basic calculation model, which has been built in matlab. this model in its first version did not apply the contact between wheels and rails. it was necessary to find out, if this simplification is relevant with respect to subsequent simulations within the complex simulation model and its results. therefore, the contact interaction as a traction force in longitudinal direction in the wheel-rail contact was realized via the kalker’s linear theory. this article deals with the comparison between models with and without the implementation of the wheel-rail contact and its influence on the damping within the torsion system and retuning of the torsion system. keywords: kalker, natural frequency, railway vehicle, torsion system, wheel-rail contact. 1. introduction one of the typical features of modern electrical railway vehicles is the individual mechanical drive of the wheelset, e.g. see figure 1. in principal it means that each wheel-set has its own propelling unit. this unit mostly consists of an electrical traction motor, a gearbox and a coupling connecting both of them. for sure there are more concepts of such a mechanical drive unit. figure 1. a partly-suspended drive of a locomotive [1]. for purposes of this research a fully-suspended drive layout has been applied. this type of an individual drive consists of an electrical traction motor, a gearbox and a hollow shaft, which transmits driving and braking power between the gearbox and the wheel-set – figure 2. because generally this research is focused on high-speed and high-power railway vehicles, the fully-suspended type of a drive train was chosen, as it is a typical and an appropriate layout for this type of vehicles. figure 2. a fully-suspended drive of a locomotive [2]. 2. basic mathematical model this mathematical model, built in matlab, is utilized to provide basic characteristics of a torsion system, which schematically represents the design of a mechanical drive train of a railway vehicle. the basic characteristics are: • natural frequencies of oscillations • natural modes of oscillations 1 https://doi.org/10.14311/app.2022.35.0001 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vojtěch dybala acta polytechnica ctu proceedings knowledge of these characteristics is supposed to be important for evaluation of frequency analysis, which will be carried out in subsequent simulations in the above-mentioned complex simulation model, built in matlab simulink. now the complex simulation model applies some simplifications to do simulations effective in terms of simulation time, amount of data and goals of the research. the basic calculation model supposes also some simplifications as the interaction in the wheel-rail contact. because of the frequency analysis it is appropriate to do a review if the simplification of the basic calculation model is reasonable and if there can be a significant impact on the evaluation frequency analysis. 2.1. model with no wheel-rail contact the mathematical model is based on the scheme of the torsion system (figure 3), which respects the layout of the fully-suspended drive presented in figure 2. figure 3. torsion system scheme – fully-suspended drive [3]. equations of motion describing this system were derived via the lagrange method. the matrix notation of these equation of motion is (1). to calculate natural modes and natural frequencies of oscillations these equations are solved as a system without the vector of excitation [m ], see (2). [j ][φ̈] + [k][φ] = [m ] (1) [j ][φ̈] + [k][φ] = [0] (2) the system (2) was solved via the function eig(k, j ), which is a pre-programed function in matlab. this function returns: • an eigen vector representing rotation angle φi,j for a rotation mass ji and j natural mode of oscillation • an eigen values vector [λj ] representing j natural angle frequency ωj = √ λj the eigen values are transformed into natural frequencies fj via the formula (3). fj = √ λj 2π (3) 2.2. model with a wheel-rail contact the wheel-rail contact was implemented into the calculation model via longitudinal traction forces t1, which can be seen in the top view of the drive train – figure 4. the wheel-rail contact is simplified so, that lateral forces and spin moment between the wheel and the rail are neglected. practically it means that it represents rolling of a cylindrical wheel on rails, not conical profile wheels and sinus movements of the wheel-set is neglected. also forces t1 on both sides are supposed to be same. figure 4. visualization of wheel-rail forces – top view. the longitudinal traction force t1 is calculated via kalker’s linear theory for the wheel-rail contact (4). t1 = c11 ael bel g sx = c1 sx = k1 sx (4) in this article details about the kalker’s theory will not be presented, but they can be found in [4– 6]. for purposes of this contribution the value of kalker’s coefficient c11 = 4, 984 is taken as a fact. that coefficient itself can varies. because that theory is linear, its validity is limited for small values of slip with respect to adhesion characteristic – phase i in figure 5. for the area of higher slip (phase ii) the direction of the curve k1* decreases. while the value of c11 can be supposed for good conditions in the wheel-rail contact (dry rails), for worst conditions (wet or dirty rails) it can decrease as well. this fact is presented in figure 6 within popovici’s adhesion characteristics. and therefore because of both effects, the calculations were carried out also for variations of c11, specifically for c11/2 and c11/4. sx = rk ωk − vx v (5) due to above mentioned simplifications also only the wheel slip sx in the longitudinal direction is supposed. its deduction is presented in figure 7 and its mathematical representation is formula (5). [j ][φ̈] + [b][φ̇] + [k][φ] = [0] (6) within the torsion system scheme (figure 8) and subsequently the equations of motion in the matrix representation (6) the force t1 creates a damping element in the wheel-rail contact. this damping defines 2 vol. 35/2022 kalker’s coefficient c11 and its influence figure 5. traction characteristic and adhesion characteristic of wheelset [4–6]. figure 6. popovici’s adhesion characteristics [7]. the formula (7). rk is constant, c1 = k1 (figure 5 and figure 6) will varies according to c11 and the vehicle velocity v is a variable as well. the calculations for this system were carried out as an analysis of the influence of c1 and v on the damping factor. bw −r = c1 r2k v (7) the velocity as a calculation parameter was specified based on the supposed traction characteristics (figure 9) of the vehicle, which is a high-speed locomotive for purposes of the research. the applied values of velocity were from 0 km/h to 200 km/h. the system (6) was solved by the matlab funcfigure 7. wheel slip deduction [2]. figure 8. torsion system scheme – fully-suspended drive. figure 9. vehicle traction characteristic. tion polyeig(k, b, m ), which returns eigenvectors and eigenvalues for defined matrices. eigen vectors returns values of angle of rotation as described in 2.1. eigen values are complex numbers in the form according to (8) for this system with damping. λj = −δj ± i ωdmp,j (8) resulting natural frequencies of oscillation are then calculated according to (9). fdmp,j = ωdmp,j 2π (9) 3 vojtěch dybala acta polytechnica ctu proceedings 3. calculation results results of calculations in sections 3.1 and 3.2 below are respective to the theory from sections 2.1 and 2.2 and the subsequent conclusion aims to make an analysis of the velocity influence on the retuning of the torsion system via damping. 3.1. model with no wheel-rail contact the torsion system scheme, which respects the design of the fully-suspended drive in figure 2 is a system with 7 degrees of freedom and seven natural frequencies of oscillation were calculated, see table 1. because the torsion system in this state was considered as a free system the first natural frequency is 0hz and it matches with a free rotation of the system. natural modes of oscillation complying with the natural frequencies are presented in figure 10 to figure 16. such a considered state of the torsion system can be agreed with the situation when a railway vehicle does not generate any force in the wheel-rail contact – standstill of the vehicle or drive without traction or brake force. natural frequencies of torsion system [hz] 1. 2. 3. 4. 5. 6. 7. 0 6 57 337 572 857 2403 table 1. natural frequencies overview [3]. figure 10. first natural mode of torsion oscillations [3]. figure 11. second natural mode of torsion oscillations [3]. table 2 provides a description of these natural modes with respect to dominant oscillation of a specific rotation mass. order respective dominant less of natural oscillations significant natural frequency of a mass oscillations modes [hz] 1. 0 own free – rotation 2. 6 wheel-set towards – hollow shaft 3. 57 wheels of – wheel-set 4. 337 wheel-set pinion towards towards hollow shaft rotor 5. 572 pinion towards – rotor 6. 857 wheel-set towards hollow hollow shaft shaft joints gear wheel towards hollow shaft 7. 2403 pinion towards rotor – pinion towards gear wheel table 2. description of natural modes [3]. figure 12. third natural mode of torsion oscillations [3]. figure 13. fourth natural mode of torsion oscillations [3]. 4 vol. 35/2022 kalker’s coefficient c11 and its influence figure 14. fifth natural mode of torsion oscillations [3]. figure 15. sixth natural mode of torsion oscillations [3]. figure 16. seventh natural mode of torsion oscillations [3]. 3.2. model with a wheel-rail contact in this section the attention will be put on calculated eigenvalues and respective natural frequencies. the first group of calculations was done for the value of the kalker’s coefficient c11 = c1. with respect to the damping figure 17 shows that, the damping reaches very high values in low velocities and strongly decrease with increasing velocity. this fact can be concluded with respect to the third eigenvalue, which relates to torsion oscillations of the wheel-set, see table 2. regarding natural frequencies see figure 18. a small change in the value of the second one can be observed, from 6hz to 4hz. a significant change can be observed regarding the third one when its value decreased from 57hz to 0hz. this means that the damping in the wheel-rail contact is so high, that it suppresses torsion oscillations of the wheel-set. for the second group of calculations, where c1 = c11/2 = k1∗, see figure 5 and figure 6, the situation changed. the damping (figure 19) was very high in low velocities again and suppressed oscillations of the wheel-set on the frequency of 57hz. with increasing velocity, the damping decreases. between 100 km/h and 105 km/h retuning of the system appeared. there figure 17. damping as a function of velocity – c11. figure 18. natural frequencies as a function of velocity – c11. is a visible jump in damping and change in the third natural frequency (figure 20), which started to grow. then the third natural frequency increases with decreasing damping and approximate to the value of 57 hz. also, a small change in the value of the second natural frequency from 6hz to 4hz was observed. figure 19. damping as a function of velocity – c11/2. in the third group of calculations, where c1 = c11/4 = k1∗, see figure 5 and figure 6, the results were equivalent to results from the second group. the difference, which figure 21 and figure 22 presents, is that the point of torsion system retuning occurs in lower velocity, between 50 km/h and 55 km/h. a small 5 vojtěch dybala acta polytechnica ctu proceedings figure 20. natural frequencies as a function of velocity – c11/2. change in the value of the second natural frequency from 6hz to 4hz was observed again. figure 21. damping as a function of velocity – c11/4. figure 22. natural frequencies as a function of velocity – c11/4. the fourth, the fifth, the sixth and the seventh eigen value and respective damping parameters and natural frequencies are not mentioned, because the calculations did not show any influence of the variation of values of the velocity and the kalker’s coefficient. 4. conclusions results of calculations presented in figure 17 to figure 22 proved, that the wheel-rail contact can significantly influence a behavior of a torsion system, which means to change some natural frequencies. specifically, the third natural frequency related to the oscillations of the wheel-set itself in a very strong way and weakly the second natural frequency related to oscillations of the wheel-set towards the hollow shaft. on the other hand, it was presented, that the variability of parameters, which characterize the wheel-rail contact, don’t influence the rest of the torsion system. with respect to a research it can practically mean, that for a research oriented on a wheel-set torsion oscillations and related problematics, the effect of the wheel-rail contact should be considered. on the other hand, within a research of torsion phenomenon regarding the rest of a torsion system, as traction motor, gears and coupling, the wheel-rail contact can be neglected. list of symbols ael main half-axis of contact ellipse [m] bel secondary half-axis of contact ellipse [m] b, bw −r damping parameter [nms.rad−1] c11 kalker’s coefficient [-] f natural frequency [hz] fdmp natural frequency of damped system [hz] j mass of rotation [kg.m2] k torsion stiffness [nm.rad−1] m external torque [nm] t1 tangential force [n] rk wheel radius [m] sx wheel slip [%] φ angle rotation [rad] φ̇ time derivative of angle rotation [rad.s−1] φ̈ second time derivative of angle rotation [rad.s−2] λ eigenvalue [-] ω natural angle frequency [rad.s−1] ωdmp natural angle frequency of damped system [rad.s−1] ωk angular speed of a wheel [rad/s] δ oscillation damping [rad/s] acknowledgements this research has been realized using the support of the technology agency of the czech republic, programme national competence centres, project #tn01000026 josef bozek national center of competence for surface transport vehicles this support is gratefully acknowledged. references [1] t. fridrichovský. studie disertační práce. praha, 2017. [2] v. dybala, m. libenský, b. šulc, c. oswald. slip and adhesion in a railway wheelset simulink model proposed for detection driving conditions via neural networks. in sborník vědeckých prací vysoké školy báňské – technické univerzity ostrava, řada strojní, ostrava, 2018. [3] v. dybala. the electromagnetically excited resonance of the pinion in fully-suspended drive of a locomotive and its sensitivity on the torsion stiffness of the rotor shaft. in sborník abstraktů konference stč, praha, 2021. 6 vol. 35/2022 kalker’s coefficient c11 and its influence [4] j. kolář. teoretické základy konstrukce kolejových vozidel. praha: česká technika – nakladatelství čvut, 2009, p. 276. [5] j. kolář. zborník prednášok ii. – xx. medzinárodná konferencia – súčasné problémy v kolajových vozidlách. in dynamika individuálního pohonu dojkolí s nápravovou převodovkou, žilina, 2011. [6] j. kolář. problémy modelování vlivu svislých nerovností trati do dynamiky pohonu dvojkolí. in železničná doprava a logistika xi, pp. 38–47, 2015. [7] r. popovici. friction in wheel-rail contacts. enschede, the netherlands: university of twente, 2010. 7 acta polytechnica ctu proceedings 35:1–7, 2022 1 introduction 2 basic mathematical model 2.1 model with no wheel-rail contact 2.2 model with a wheel-rail contact 3 calculation results 3.1 model with no wheel-rail contact 3.2 model with a wheel-rail contact 4 cocnlusions list of symbols acknowledgements references https://doi.org/10.14311/app.2022.33.0452 acta polytechnica ctu proceedings 33:452–456, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague dowels effectiveness investigation between ultra high performance fiber reinforced concrete and reinforced concrete spyridon a. paschalisa, ∗, andreas p. lampropoulosb a school of engineering and computing, university of west london, uk b school of architecture, technology and engineering, university of brighton, uk ∗ corresponding author: spyros.paschalis@uwl.ac.uk abstract. in the present study, the performance of reinforced concrete (rc) beams which were strengthened with ultra high performance fiber reinforced concrete (uhpfrc) and dowels at the interface was investigated. rc beams with a length of 2.2 m strengthened with uhpfrc layers at the tensile side. before the application of the layers, the interface between rc and uhpfrc was roughened. during the testing, the interface slips between uhpfrc and rc were recorded using linear variable differential transformers (lvdts). the beams were tested under four-point flexural test. the results of the present study indicated that the dowels at the interface reduce the slips at the interface, delay the formation of cracks and result in higher load carrying capacity. keywords: dowels, interface characteristics, reinforced concrete beams, steel connectors, strengthening, uhpfrc. 1. introduction safety of structures is of paramount importance especially in case of old substandard structures or structures damaged by earthquakes or other accidental actions. existing strengthening and retrofitting methods present crucial disadvantages which are related to the high cost, the difficulty in the application and the low performance of the technique. the present study investigates the effectiveness of ultra high performance fiber reinforced concrete (uhpfrc) as strengthening material using dowels at the interface between reinforced concrete (rc) members and uhpfrc layers. the are several studies in the literature investigating the properties of the uhpfrc [1–4]. the last years research has been focused on the application of uhpfrc as repair and strengthening material. an analytical model for the prediction of the response of composite uhpfrc and rc elements was developed by habel et al. [5]. uhpfrc was applied as repair material from bruhwiler and denarie [6], while lampropoulos et al [7] developed a numerical model that predicts the performance of rc beams which were strengthened with uhpfrc. paschalis et al. [8], conducted full scale tests to investigate the effectiveness of uhpfrc for the strengthening of rc beams. in this study, rc beams strengthened with uhpfrc layers, with and without the use of steel bars in the layers. the addition of steel bars in the layer resulted in high load carrying capacity of the strengthened beams. however, high values of slips at the interface between uhpfrc and concrete were recorded. in the present study the performance of rc beams which were strengthened with uhpfrc layers and dowels at the interface has been investigated. this is a topic which has not been investigated in detail and needs further investigation. 2. experimental program six beams have been tested in the present study. two rc beams prepared without any layer and used as control beams. in two beams, uhpfrc layers added at the tensile side and in two beam dowels added at the interface between uhpfrc and rc. the reinforcement and the dimensions of the beams without any reinforcement are presented in figure 1. in figure 2, the dimensions of the layers and the position of the dowels are presented. the dowels were designed based on greek retrofitting code [9]. 3. preparation of the specimens fine sand, ggbs, silica fume, superplasticizer, high strength cement and steel microfibers were used in the mix design. in addition, a high percentage of fibers (3 vol.%) was incorporated in the mix to achieve strain hardening. the interface was roughened using a needle scaler and the dowels placed in position using a drill. the preparation of the beams which were strengthened with dowels is presented in figure 3. 4. properties of the materials cubes with 10 cm sides, tested in compression according to bs en 12390-3:2009 [10]. the compressive 452 https://doi.org/10.14311/app.2022.33.0452 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 33/2022 dowels effectiveness investigation figure 1. control beam. figure 2. strengthened beams using dowels. figure 3. preparation of the beams. strength of concrete which was used for the preparation of the beams was 31 mpa and the compressive strength of uhpfrc layers was 136 mpa. the tensile characteristics of uhpfrc identified with a direct tensile test. the results are presented in figure 4. based on the average curve of figure 4, the tensile strength was 11.5 mpa and the modulus of elasticity 50 gpa. 5. testing of the beams in figure 5, a strengthened beam with layer before the testing is presented. as can be seen, lvdts were place at the interface between uhpfrc and rc to record the slip. in figure 7, the average experiments results are placed. the results of figure 7 indicate the effectiveness of dowels. in both techniques, the stiffness of the beams was increased. however, with the addition of dowels a higher load carrying capacity was achieved. in this case, the maximum load was increased by 22.5%. on the contrary, an increase of only 1.5% was achieved with the addition of layers only. in figure 8, the beams after the testing are presented. the control beams failed with a major crack at the tensile side (figure 8a). the failure mode of the two beams which were strengthened with uhpfrc 453 spyridon a. paschalis, andreas p. lampropoulos acta polytechnica ctu proceedings figure 4. stress-strain curve from the testing of uhpfrc in tension. figure 5. experimental setup for the testing of the beams. figure 6. a) lvdts on side 1 b) lvdts on side 2. 454 vol. 33/2022 dowels effectiveness investigation figure 7. results from the testing of beams. figure 8. failure mode a) control beam b) beam strengthened with layer c) beam strengthened with layer and dowels. layers was not the same. in one beam the bond at the interface was strong and a major crack at the layer progressively propagated to the rc beam. in the second beam, a de-bonding at the interface occurred (figure 8b). finally, identical was the failure mode of beams which were strengthened with layers and dowels. a major crack appeared in the layers and propagated to the rc beam (figure 8c). during the testing it was noticed that dowels resulted in the delay of the formation of cracks. this was also evident from the results of figure 7. the cracking on the control beams started at value of load 5 kn, on the strengthened beams with layers at 15 kn and on beams with layers and dowels at 24 kn. the results from the measurements of slips at the interface are presented in table 1. due to the fact that a local debonding commenced in one of the beam which was strengthened with uhpfrc layer, the reading of this beam were ignored. lvdt layers 1 dowels 1 dowels 2 2 0.06 0.04 0.01 3 0.11 0.01 0.18 4 0.36 0.05 0.06 7 0.03 0.03 0 8 0.18 0.14 0.01 table 1. reading of the lvdts. from the results of table 1, it can be noticed that almost in all positions, the values of slip for the beams with dowels at the interface, were significantly lower 455 spyridon a. paschalis, andreas p. lampropoulos acta polytechnica ctu proceedings compared to the slips of the beam without dowels. this indicates the better bonding at the interface when dowels are used. 6. conclusions the main aim of the present study was to investigate the effectiveness of dowels at the interface between uhpfrc and rc. full scale tests were conducted and rc beam with a length of 2.2 m strengthened with uhpfrc layers, with and without dowels. the slips at the interface were also recorded. based on the experimental results the following conclusions can be drawn: • dowels delay the formation of the cracking • dowels result in a better bonding • dowels result in a higher load carrying capacity based on the findings of the present research, the use of dowels should be considered in cases that better bonding and better performance of the strengthened elements is required. references [1] l. ferrara, n. ozyurt, m. di prisco. high mechanical performance of fibre reinforced cementitious composites: the role of casting-flow induced fibre orientation. materials and structures 44(1):109-28, 2010. https://doi.org/10.1617/s11527-010-9613-9. [2] s. paschalis, a. lampropoulos. fiber content and curing time effect on the tensile characteristics of ultra high performance fiber reinforced concrete. structural concrete 18(4):577-88, 2017. https://doi.org/10.1002/suco.201600075. [3] s. paschalis, a. lampropoulos. ultra high performance fiber reinforced concrete under cyclic loading. aci materials journal 113(4):419-427, 2016. [4] s.-t. kang, j.-k. kim. the relation between fiber orientation and tensile behavior in an ultra high performance fiber reinforced cementitious composites (uhpfrcc). cement and concrete research 41(10):1001-14, 2011. https: //doi.org/10.1016/j.cemconres.2011.05.009. [5] k. habel, e. denarié, e. brühwiler. structural response of elements combining ultrahigh-performance fiber-reinforced concretes and reinforced concrete. journal of structural engineering 132(11):1793-800, 2006. https://doi.or g/10.1061/(asce)0733-9445(2006)132:11(1793). [6] e. bruhwiler, e. denarie. rehabilitation of concrete structures using ultra-high performance fibre reinforced concrete the second international symposium on ultra high performance kassel, germany, 2008. https://infoscience.epfl.ch/reco rd/182195/files/010.pdf. [7] a. p. lampropoulos, s. a. paschalis, o. t. tsioulou, et al. strengthening of reinforced concrete beams using ultra high performance fibre reinforced concrete (uhpfrc). engineering structures 106:370-84, 2016. https: //doi.org/10.1016/j.engstruct.2015.10.042. [8] s. a. paschalis, a. p. lampropoulos, o. tsioulou. experimental and numerical study of the performance of ultra high performance fiber reinforced concrete for the flexural strengthening of full scale reinforced concrete members. construction and building materials 186:351-66, 2018. https: //doi.org/10.1016/j.conbuildmat.2018.07.123. [9] greek organization for seismic planning and protection, greek retrofitting code (greco), athens, 2013. [10] bs en 12390-3:2009. testing hardened concretepart 3: compressive strength of test specimens, 2009. 456 https://doi.org/10.1617/s11527-010-9613-9 https://doi.org/10.1002/suco.201600075 https://doi.org/10.1016/j.cemconres.2011.05.009 https://doi.org/10.1061/(asce)0733-9445(2006)132:11(1793) https://infoscience.epfl.ch/record/182195/files/010.pdf https://doi.org/10.1016/j.engstruct.2015.10.042 https://doi.org/10.1016/j.conbuildmat.2018.07.123 acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0424 acta polytechnica ctu proceedings 38:424–428, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague bringing students closer to the issues of indoor environmental quality and technology using nemos device: an experience in albania francesco salamonea, ∗, lorenzo belussia, ludovico danzaa, etleva dobjanib, matteo ghellerea, italo meronia, arben shtyllab, saimir shtyllab a itc-cnr, via lombardia 49, 20098 san giuliano milanese, italy b polis university, department of applied research, rr. bylis 12, autostrada tiranë-durrës, km 5, 1051, tirana, albania ∗ corresponding author: salamone@itc.cnr.it abstract. the article presents the outcomes of a monitoring campaign and a survey performed in a school building in tirana. included in the bilateral agreement between the national research council of italy and the ministry of education and sport of the republic of albania, the research is based on environmental analysis, collection of data from monitoring of environmental variables and students’ feedback. the monitoring of indoor environmental quality (ieq) was carried out with a nearable environmental monitoring system (nemos) for ieq purposes, designed and developed by itc-cnr based on the do-it-yourself (diy) philosophy. the proliferation of the maker movement philosophy has promoted the diffusion of diy-based technologies. the spread of this movement is observed outside schools but there is a growing interest among educators to introduce this philosophy in the students’ education, approaching to steam (science technology engineering arts mathematics) in a different and more stimulating way. the application of two nemos devices in two classrooms (in different locations) aroused the curiosity of students, bringing them closer to the world of technology with a practical example. it also increased their awareness about the overall ieq conditions in the classrooms where they spend a considerable part of their lives. keywords: ieq, albania, school, sustainability, survey. 1. introduction school buildings are fundamental in the education of future generations and to transmit them the desire for research and innovation, too. coding and programming have emerged as the most desired success factors for society [1] where especially for the first one there should be considered its introduction in the didactic curricula in response to the labour market demand. the main difference among them is that coding is a part of programming and involving the translation of a written code by a machine; while the second is the process of creating a program following specific standards and performing a specific task, too. in the last decade, researchers and educators have developed tools to simplify the mechanics of learning computing at school, introducing new programmers to become more fluent and expressive towards new technologies [1]. new technological developments of microcontrollers, low-cost sensors and actuators foster the teaching of these new concepts [2]. countries that are distinguished for their technologically advanced economies and societies are reconsidering their education policies by concentrating their economic efforts in the promotion of the stem education (science, technology, engineering and mathematics) [3]. in this sense, arduino-based projects can stimulate childrens’ computing competences [4] as well 3d-printing approach can be a promising technology to engage students with hands-on interactions [5]. in the last years, the indoor environmental quality (ieq) in schools has attracted the attention of several researchers, since the quality of indoor spaces affects children’s well-being, health, and productivity. scientific literature reports several case studies focused on the evaluation of the ieq in educational buildings, where children and teachers were inquired about their perception of the indoor environment. as planned in the two-year bilateral agreement between the national research council of italy and the ministry of education and sport of the republic of albania, a monitoring campaign was designed and conducted in a school in the outskirts of tirana but in a reduced number of effective days due to covid19. therefore, it can be considered as a small-scale test and a useful reference to be reproduced in similar scenarios with a longer time span. the short three-day monitoring campaign made possible to comprehend whether the methodology based on the use of the nemos device [2], developed in the laboratories of itc-cnr, can be used, on the one hand, to collect data on subjective and objective environmental quality in the classrooms 424 https://doi.org/10.14311/app.2022.38.0424 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 bringing students closer to the issues of indoor environmental quality . . . figure 1. school building with sun path for the considered class period (november 16th–18th), classrooms configurations and position of nemos devices (in red) during the campaign. where students spend most of their lives; on the other hand, to raise students’ awareness regarding the importance of assessing the environmental quality in the classrooms thanks to a diy-based and 3d-printed monitoring device. in this case, the pupils are not involved in the production of the devices, but they could be in the future. this pilot project can be seen as an opportunity by stakeholders to integrate new technical disciplines in the educational background. to acknowledge pupils and teachers with the results of the experiment, the research team delivered a final report to the school’s board. 2. case study and methodology the case study is the “kolë jakova” school (figure 1), located in the north-western suburb area of tirana (lat. 41° 21′ 28′′ n, long. 19° 45′ 44′′ e). built in 19831 and with a capacity of 12-classrooms2, it is among the most diffused typologies in rural areas during the period 1945–1990 in albania. nowadays its capacity represents a limit because the demand for classrooms is higher and consequently it constraints the split of the didactic activity for some classes in the morning and for others in the afternoon. indeed, the school hosts 420 pupils for a total of 18 classrooms (elementary and lower secondary cycle) but during the monitoring campaign all classrooms were at half capacity due to the covid-19 restrictions. the classrooms chosen for the campaign were positioned on the opposite sides of the building (east and 1referring site plan file at the central national technical archive: “shkollë në fshatin bregu i lumit dt 10.05.1983” (school in the village of bregu i lumit dated 10.05.1983). 2technical name of the project at the central national technical archive: “shkollë 8-vjeçare me 12 dhe 8 klasa 73.6” (8-year cycle school with 12 and 8 classrooms 73.6). west facing) to comprehend the impact of different orientations on people comfort perception. besides being built in the 80s it has undergone to partial renovations including the installation of a heating system not expected in the original design. the vertical opaque components are uninsulated brickbased walls (u-value = 1.26 w/m2k). the flat roof was insulated with 0.05 m thick polystyrene layer with a waterproof layer on top (u-value = 0.48 w/m2k). the windows (130 × 160 cm) are single-glazed with aluminium frame (u-value of 6.8 w/m2k and shgc = 0.7). the heating system is a centralized two-stage plant with power ranging between 178–391 kw with radiators in each classroom, but inactive during the survey, whereas two separate boilers guarantee the production of domestic hot water (dhw). the monitoring campaign and the survey aiming to collect of subjective and objective data was carried out between november 16th–18th, 2020 in the middle fall/winter school session in albania. the objective monitoring of indoor environmental variables was carried out using nemos [3, 4]. before the campaign, students were introduced to concepts of computing and 3d-printing (the pillars nemos was built) and ieq. the nemos device gained great attention among students in both classrooms. they were not informed regarding any instant value of the acquired data to avoid any eventual influence on their responses. before the start of the campaign, students were informed regarding the purpose of the monitored environmental factors (efs) measured by the nemos device. the monitoring campaign was performed for three consecutive days in each classroom. a simplified survey form was elaborated to acquire the subjective data of pupils’ perceptions in order to avoid any eventual confusion and misunderstanding. as illustrated in the figure 2, the questions were focused on the thermal 425 f. salamone, l. belussi, l. danza et al. acta polytechnica ctu proceedings figure 2. questionnaire structure about thermal, visual and iaq perception (translated from albanian). classroom total [-] age [y] weight [kg] height [cm] iclo [clo] metst [met]total (m–f) avg ± std avg ± std avg ± std avg ± std avg ± std east 13 (6–7) 14 ± 0.5 57.92 ± 14.38 164.32 ± 13.01 0.73 ± 0.11 1 ± 0 west 8 (4–4) 13 ± 0.5 50.42 ± 6.42 166.73 ± 6.62 0.78 ± 0.10 1 ± 0 table 1. data of the 21 students involved in the study. and visual perception, based on the five-point bipolar scale (from “cool” or “dark” = −2 to “warm” or “bright” = +2), while the question about the indoor air quality (iaq) perception was based on a unipolar scale (from “not smelly” = 0 to “very smelly” = 3). as mentioned above, due to the restriction because of covid-19, the classrooms were not at full capacity. the subjective data were collected from the pupils’ perceptions expressed in the specific questionnaire regarding the three parameters of the ieq during the didactic activity. during three consecutive one-hour intervals, between 1.00 p.m. to 3.00 p.m. the pupils had to express in the questionnaire their thermal, visual and iaq perceptions. all the 21 pupils that took part in the campaign, have been grouped for each class in the aggregated data as listed in table 1. students brought the parental consent to take part in the survey. from the analysis of participants’ answers it was possible to calculate the thermal resistance of their clothing in compliance with annex c of the standard en iso 7730. an additional thermal resistance of 0.1 clo for sedentary activities due to the standard school chair is considered [5]. the standard metabolic rate metst was defined in accordance with the value reported in annex b of en iso 7730. 3. results and discussions for a better evaluation of the impact of the orientation and direct daylight exposure, the chosen classrooms are positioned on the opposite sides of the building as illustrated in figure 1. in figure 3, there are reported the environmental data acquired during the monitoring campaign in the chosen classrooms, whereas in figure 4 are reported the subjective responses of students on the considered aspects. the assessed classrooms have similarities regarding the measured levels of air velocity (av) and relative humidity (rh). the av levels highlight lower averages and limited variation, determining ideal conditions for low draught. on the other hand, rh levels highlight moderate variations and mean values slightly higher than 60 %, not distant from the standard design value of 50 %. there is a slight difference regarding the illumination (il) levels between the classrooms: although the mean il values are almost the same, the data variation in west classroom (yellow device) is higher compared to the east classroom (green device). evident differences between the two classrooms emerge from the comparison of the operative temperature (ot) and carbon dioxide concentration (co2) levels in terms of mean values and variability. indeed, in the east classroom (green device) are registered the highest mean values for ot (19.8 °c vs 18.5 °c) and co2 (1800 vs 1200), as well the highest co2 variability, while the west classroom (yellow device) highlights the highest ot variability. the huge variability in co2 concentration is essentially due to the natural ventilation assured in both classrooms [6]. the analysis of subjective responses partially confirms the ieq monitored data (figure 4). analysing visual comfort (vc) perception, 74 % and 71 % of the respondents respectively report a neutral room brightness. in east classroom the percentage of pupils perceiving a “slightly bright” condition is more than three time higher than in the west classroom (18 % and 5 % respectively). and conversely the percentage of pupils perceiving the “slightly dark” condition is three time higher in west than xib (24 % and 8 % respectively). previous results suggest that a mean 426 vol. 38/2022 bringing students closer to the issues of indoor environmental quality . . . figure 3. monitored environmental data: a) east classroom; b) west classroom. figure 4. subjective perception of thermal comfort (tc), visual comfort (vc) and indoor air quality (iaq): a) east classroom; b) west classroom. illuminance value around 300 lx implies a perceived neutral or bright environment for at least 3 out of 4 users, confirming the findings of [7] for similar buildings and students age. the iaq perception partially reflects the analysis of the corresponding monitored data (co2). despite a consistent difference in co2 concentration values (the east value is 600 ppm higher than the west value), only 7 % of east perceive the air as “very smelly”, while the most common perception for both classrooms is “1 – low smelly”. a reason for this could be the low sensitivity of users to air quality, since other ieq aspects are more immediate to assess (i.e., room brightness, temperature level). even though the reduced number of pupils for classrooms, due to the pandemic, the finding related to iaq data, both monitored and perceived, are in line with [8]. finally, the tc analysis confirms the measured data (east classroom warmer than west classroom). the east classroom shows a higher neutral perception (62 % vs 48 %) and a lower “−2 – cool” perception (0 % vs 14 %) than the west classroom. the tc subjective responses also reflect the main ot variability in west classroom as long as it causes a high thermal adaptation body activity of the users. 4. conclusions the monitoring campaign and the survey have been performed in an unusual period because of the ongoing pandemic situation worldwide with didactic activities performed with reduced number of pupils. another important limitation of the study is related to the resolution of the monitoring campaign: only one measuring point for each classroom. in addition, it was not possible to record the status of opening or closing doors and windows. this would have allowed identification, without speculation, of what the effect on co2 concentrations might be. this is an improvement that could be considered in a future monitoring campaign. nevertheless, monitoring based on the use of low-cost, 3d-printed solutions represents a great stimulus to arouse students’ interest in the topic of growing field such as (iot) and increases their awareness regarding the indoor comfort issues. the main differences of monitored data between the two classrooms are identified in ot with the lowest mean value recorded in the west classroom, and co2, with the highest value monitored in east classroom. probably the main differences in ot data can have influenced the temperature comfort (tc) perception, thus allowing to justify 14 % of west classroom perceiving “cool”. the 427 f. salamone, l. belussi, l. danza et al. acta polytechnica ctu proceedings co2 differences are not reflected by iaq perception and this probably may reveal students’ adaptation to this environment condition [7]. these aspects could be considered to improve the overall ieq in classrooms performing participatory activities where students can be engaged with hands-on interactions, using self-developed monitoring devices. in this sense, it might be possible to bring students close to the themes of ieq and technology. measurement performed on a large scale, with devices such as nemos, along with feedback provided by a larger number of participants, will make it possible to identify more accurate differences between objective data and subjective perceptions in relation to the considered efs and the target audience. acknowledgements the authors thank the ministry of education and sport of the republic of albania for the authorization to conduct the monitoring campaign. references [1] p. tuomi, j. multisilta, p. saarikoski, j. suominen. coding skills as a success factor for a society. education and information technologies 23:419–434, 2018. https://doi.org/10.1007/s10639-017-9611-4 [2] f. salamone, l. danza, i. meroni, m. c. pollastro. a low-cost environmental monitoring system: how to prevent systematic errors in the design phase through the combined use of additive manufacturing and thermographic techniques. sensors 17(4):828, 2017. https://doi.org/10.3390/s17040828 [3] f. salamone, l. belussi, l. danza, et al. design and development of nemos, an all-in-one, low-cost, web-connected and 3d-printed device for environmental analysis. sensors 15(6):13012–13027, 2015. https://doi.org/10.3390/s150613012 [4] f. salamone, l. danza, i. meroni, m. c. pollastro. a low-cost environmental monitoring system: how to prevent systematic errors in the design phase through the combined use of additive manufacturing and thermographic techniques. sensors 17(4):828, 2017. https://doi.org/10.3390/s17040828 [5] t. wu, w. cui, b. cao, et al. measurements of the additional thermal insulation of aircraft seat with clothing ensembles of different seasons. building and environment 108:23–29, 2016. https://doi.org/10.1016/j.buildenv.2016.08.008 [6] details about natural and mechanical ventilation. [2022-03-21]. https://www.epfl.ch/labs/hobel/mechanical-ornatural-ventilation/ [7] l. noda, a. v. p. lima, j. f. souza, et al. thermal and visual comfort of schoolchildren in air-conditioned classrooms in hot and humid climates. building and environment 182:107156, 2020. https://doi.org/10.1016/j.buildenv.2020.107156 [8] l. dias pereira, l. neto, h. bernardo, m. gameiro da silva. an integrated approach on energy consumption and indoor environmental quality performance in six portuguese secondary schools. energy research & social science 32:23–43, 2017. https://doi.org/10.1016/j.erss.2017.02.004 428 https://doi.org/10.1007/s10639-017-9611-4 https://doi.org/10.3390/s17040828 https://doi.org/10.3390/s150613012 https://doi.org/10.3390/s17040828 https://doi.org/10.1016/j.buildenv.2016.08.008 https://www.epfl.ch/labs/hobel/mechanical-or-natural-ventilation/ https://www.epfl.ch/labs/hobel/mechanical-or-natural-ventilation/ https://doi.org/10.1016/j.buildenv.2020.107156 https://doi.org/10.1016/j.erss.2017.02.004 acta polytechnica ctu proceedings 38:424–428, 2022 1 introduction 2 case study and methodology 3 results and discussions 4 conclusions acknowledgements references my title https://doi.org/10.14311/app.2022.36.0001 acta polytechnica ctu proceedings 36:1–5, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague reliability of a structure stricken by a tornado milan holický czech technical university in prague, klokner institute, šolínova 7, 166 08 prague, czech republic correspondence: milan.holicky@cvut.cz abstract. a tornado may become a very dangerous climatic action for any structure, depending on the tornado intensity and extent of the structure. the origin of a tornado, its range and further activity appear to be unpredictable and completely random. the strike of a tornado to a structure depends on the tornado movement and the structure extent. the tornado intensity is specified by the wind speed, following observed damages and consequences. the probability of a tornado strike on a structure depends on the extent of the structure itself, on the tornado range, and the extent of the considered reference area, for which data concerning tornado appearance are available. the probability of failure of a structure is given by the product of the probability of contact of a tornado with the structure and the probability of exceeding the design speed of the tornado wind. the probability of exceeding the design wind speed depends on the assumed distribution of wind speed and the frequency of tornadoes during the required lifetime of the structure. the target failure probability of common structures presented in available documents is 10−6 per year, and 10−7 per year of structures in power plants. however, it is not clear how these criteria have been derived. further development of reliability theory of tornado-stricken structures is recommended to be focused on the risk analysis of appropriate systems, of which the considered structures are elements, on the target failure probability, on detailed analysis of the probability distribution of tornado wind speed and their frequency during the required lifetime of the structures. keywords: design wind speed, power plant, probability, structure, target failure probability, tornado. 1. introduction reliability analysis of a structure stricken by a tornado is treated in several documents and prescriptions [1–3] that are focused primarily on nuclear power plants. recent document [2] provides detailed recommendations for reliability verification of power plants using probabilistic methods considering random properties of structures and tornado effects. all the available documents include analysis of two inseparable factors of tornado action affecting the reliability of a structure: (1) tornado strike on structure, (2) the intensity of tornado effects. both factors have random character depending on several local circumstances and conditions that are often assessed based on observed consequences and available experience. the factor (1) depends on the frequencies of the tornado in the reference territory where the tornado actions are observed. specification of such a territory depends on social and economic aspects, on longtime observation of tornado, and empirical meteorological information. tornado strike on the considered structure depends surface area of the reference territory. it is a random variable that can be only assessed. the factor (2) the intensity of tornado effects is commonly described by wind speed in the air twirl. the wind speed is again a random variable having a large variability of unknown parameters that can be assessed based on observed destructive consequences of tornado action. the design value of the wind speed can be estimated assuming usually weibull distribution and approximate parameters. because of uncertain parameters (including the skewness) the estimated design wind speed may be imprecise. the tornado intensity is usually classified using graduated tables for ground wind speeds. two scales indicated in table 1 are commonly applied: fuji and torro scales. fuji scale (having 6 grades) is primarily used in usa, torro scale (refine fuji scale having 10 grades) is used in europe. the lowest grade of both scales starts with the wind speed of 17 m/s, the highest grade from 121 m/s. 2. probability of structural failure the probability of structural failure pf due to tornado action can be assessed considering probability pa of tornado strike on a given structure, and simultaneous exceedance of wind speed of the design value vd, when the structure may fail with the probability pv . the resulting failure probability pf is therefore given as the product of two probabilities: pf = pa × pv (1) probability pa of tornado strike depends [2] on area of reference territory a0, for which records of tornado actions are available for n years, on extend of structure ak and also on the total tornado extent ator. above 1 https://doi.org/10.14311/app.2022.36.0001 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en milan holický acta polytechnica ctu proceedings fuji torro damage wind detailed description of damage f0 t0t1 slight damage 17-24 m/s 25-32 m/s accidentally demolished chimneys and wooden fences, minor damage to the roofing, damaged advertisements and road signs next to roads, broken tree branches, sporadically uprooted trees with shallow roots, traces of a tornado are visible in the field. f1 t2t3 moderate damage 33-41 m/s 42-51 m/s partially destroyed roofing, moving cars are pushed off the roads, mobile homes are shifted from the foundations, overturned or heavily damaged, flimsier buildings and shelters are badly damaged to destroyed, larger trees with stronger roots are sporadically uprooted and broken f2 t4t5 moderate damage 52-61 m/s 62-72 m/s the roofs of less rigidly built buildings are completely torn down, mobile homes and flimsier structures are completely destroyed, in brick and stronger wooden houses, the side and front walls are not yet seriously damaged. light cars float, small and light debris become projectiles, most isolated large trees are uprooted or broken. f3 t6t7 significant damage 73-83 m/s 84-95 m/s roofs and some walls are completely torn from the structure even in well-built buildings, heavier cars are buoyed, trains and locomotives are overturned, most trees in continuous forests are uprooted and broken, standing trees and tree stumps are partially debarked by flying debris. f4 t8t9 severe damage 96-107 m/s 108-120 m/s reinforced concrete buildings are significantly damaged, brick (brick) and stone buildings are severely (mostly irreparably) damaged, less solid buildings are completely levelled, the ruins of flimsy buildings are scattered at considerable distances from their foundations, cars are drifted by the air (just above the ground). ) or towed over long distances, large and heavy "projectiles" are formed from flying debris, tree stumps are completely stripped of bark. f5 t10 completedestruction > 121 m/s reinforced concrete buildings are badly damaged, other buildings are completely destroyed, cars are carried through the air like projectiles over considerable distances, fields are completely devoid of vegetation, crops are mostly torn out even with roots. table 1. tornado intensity scales of fuji and torro. mentioned variables a0, ator and ak are random variables of considerable variability that require statistical and probabilistic analysis. in addition, tornado extent ator may be assessed as product of the length l and width w (ator ≈ l × w) [2]. assuming that the tornado actions occur within the reference territory a0 uniformly, then the probability pa for one year can be estimated from the mean values of involved variables ak, ator and a0 using the formula: pa ≈ (ak + ator)/a0/n, (2) where n denotes the number of years when tornados are observed in the reference territory a0. more detailed probabilistic analysis concerning the specific location of a structure can be provided on the basis of the local conditions within the reference territory a0. an example. the probability pa of a tornado strike per year is strongly dependent on the extent of reference area a0. in south moravian region a0 = 7188 km2 and area ak + ator = ak + l × w = 1 + 4 × 0, 5 = 3 km2. then the probability pa determined from one observation is about ≈ 3/7188 ≈ 4 × 10−4. however, when a0 is considered as a circle with the centre at the structure and the radius of 100 km, then a0 ≈ 31400 km2, and pa ≈ 3/31400 ≈ 1 × 10−4. 2 vol. 36/2022 reliability of a structure stricken by a tornado design values vd [m/s] for pv parameters vinf [m/s] µ [m/s] σ [m/s] α 10−5 10−6 10−7 weibmin 30 [m/s] 42 10 1,5 127 142 157 weibmin 7 [m/s] 42 10 1,0 108 118 127 weibmin 21 [m/s] 42 10 0,5 92 98 104 weibmin 10 [m/s] 42 10 0 80 84 87 normal ± inf 42 10 0 85 90 94 table 2. design values of wind speed vd for given skewness α and exceedance probability pv . 3. design wind speed the design value vd of the wind speed v is defined for the specified exceedance probability pv when the structural failure may occur, so: pv ≈ p (v > vd), (3) weibull distribution of wind speed v is considered in available documents [1–3]. it is the lower bound limited distribution characterized by three parameters: the mean µ, standard deviation σ a and skewness α. in the professional literature and software products (mathcad, excel, easyfit) another three parameters are also used: for example c1 (also α or b), c2 (also b or a) and the lower bound vinf (for standardized variables uinf ). the parameter c1 depends solely on the skewness α, the parameter c2 depends on the parameter c1 (skewness α) and standard deviation σ, the lower bound of the standardized variable uinf depends only on the parameter c1 (on skewness α), the lower bound of the real variable vinf depends on three parameters: on the mean µ, standard deviation σ and also on the skewness α (on parameter c1) [4]. the skewness α affects therefore the lower bound vinf and shape of the distribution for a given mean and standard deviation σ, and significantly affects resulting probability of tornado effect. appendix provides a copy of mathcad sheet enabling application of the weibull distribution. an example of determining the design values of wind speed vd for given skewness α and exceedance probability pv is indicated in table 2. it follows from table 1 and figure 1 that the relatively high skewness α = 1, 5 leads to high design values vd of wind speed, by 50 % greater than for the skewness α = 0 (normal distribution). however, it is not clear how the actual skewness α can be determined from available data. it may be also recommended to utilize experience from the distribution of usual wind speed or the information concerning lower bound vinf indicated in table 1. 4. number of tornado strikes during the required lifetime of a structure several tornados, in general, say n , may occur [3]. if the distribution function of a singular tornado is φ(x, 1), then the overall distribution function of all n tornados is given by power function φ(x, 1)n [4, 5]: figure 1. weibull and normal distribution for µ = 42 m/s and σ = 10 m/s. φ(x, n ) = φ(x, 1)n (4) then the design value (upper fractile) vd,n of the distribution function φ(x, 1)n may be greater the design value determined from a single distribution φ(x, 1). the occurrence number n may be in the case of less frequent tornados equal to the number of years n used in equation 2. variation of the design value vd,n of tornados for given exceedance probabilities with the occurrence number n is indicated in table 2. probability density functions and partly distribution functions for occurrence numbers n = 1 and 10 are shown in figure 3. an example. table 3 demonstrates that the resulting design values vd,n for given number of tornados n > 1 are greater than the values for one tornado n = 1. with increasing n the design value increases up to 35 %. consequently, when designing a structure the number of tornados n during the assumed lifetime is necessary to consider. when during the assumed lifetime of a structure the number of tornados is n , then the probability pv,n , that the wind speed vn exceeds the design values vd with the probability: pv,n = p(vn > vd) (5) 3 milan holický acta polytechnica ctu proceedings occurrence number of tornados lower bound mean value standard deviation skewness design value vd,n for exceedance probability n vinf [m/s] µ [m/s] σ [m/s] α 10−5 10−6 10−7 1 30 42 10 1,5 127 142 157 10 30 61 10 1 142 157 172 100 30 80 10 1 157 172 187 1000 30 98 10 1 172 187 201 table 3. variation of the design value vd,n of tornados with the occurrence number n . figure 2. distributions of tornado wind speed for n = 1 and 10. table 4 indicates the probability p(vn > vd) for design values vd determined for a single tornado (n = 1) when structural failure may occur. design values vd are determined as before for the probabilities 10−5, 10−6 a 10−7. the probability p(vn > vd) for the design value vd determined from one observation (n = 1) is therefore increased by the order of tornado occurrence n . 5. failure probability failure probability of a structure due to tornado strike is generally given by equation1 1 as the product of the probability pa, of tornado strike, and the probability pv , that wind speed exceeds design value vd when structural failure may occur. in the case of one isolated tornado the failure probability pf,1 is given as: pf,1 = pa × pv = p(a0, ak, ator) × p(v > vd) (6) in the case of n tornados during the required lifetime the failure probability pf,n is given by the product: pf,n = pa × pv,n = p(a0, ak, ator) × p(vn > vd) (7) it is obvious that when several tornados (n > 1) are assumed, the failure probability is greater than when only one tornado (n = 1) is considered. an example. consider as before the reference area a0 = 7188 km2 and extent ak + ator = 2 + 1 = 3 km2. then the probability pa was determined from one observation ≈ 3/7188 ≈ 4 × 10−4. when the structure is design assuming the probability p(v > vd) = 10−5, then the expected probability is ≈ pf,1 ≈ pa × pv = 4 × 10−9. however, when during the assumed lifetime number of tornados is n = 10, then the failure probability is about ≈ pf,n = pa × pv,n ≈ 4 × 10−8. thus, it is the failure probability ten times greater. 6. target probability of failure target probability of failure of a structure during the lifetime t denoted ptar,t can be specified based on probabilistic optimization (maximization) of structural utility (cost) ctot(t) during its lifetime t. this can be analyzed in a simplified way considering the total benefit of functioning structure cben(t) during the lifetime t, and the possible losses cf (t) caused by structural failure (within the relevant system) with the target probability ptar,t. the total utility ctot(t) can be then symbolically expressed as [4, 5]: ctot(t) = max{cben(t) − ptar,t × cf (t)}. (8) the cost cben(t) includes all kind of benefits during the life time t. consequences of structural failure cf (t) cover all the material, ecological and social losses caused by the structural failure. these losses cf (t) may be extremely large, and could be very difficult to assess. a detailed analysis is outside this contribution and it is recommended to accept a very small target probability, for example, ptar,t = 10−7 per year as indicated in documents [1–3]. note that the target probability for common structures is usually accepted as ≈ 1.30 × 10−6 for one year (that is reliability index ≈ 4, 7). this corresponds to ≈ 0, 72 × 10−4 per 50 years (and to the reliability index 3,83), per 100 years the target probability is ≈ 1, 31 × 10−4 and reliability index ≈ 3, 65. variation of the target probabilities ptar,t with reference time t given in years for selected basic probabilities ptar,1 = 10−5, 10−6 and 10−7 is shown in figure 3. 4 vol. 36/2022 reliability of a structure stricken by a tornado number of tornados lower bound mean value standard deviation skewness the probability p(vn > vd) for design value vd n vinf [m/s] µ [m/s] σ [m/s] α 127 142 157 1 30 42 10 1,5 1.0×10−5 1.0×10−6 1.1×10−7 10 30 61 10 1 9.4×10−5 1.0×10−5 1.1×10−6 100 30 80 10 1 9.3×10−4 1.0×10−4 1.1×10−5 1000 30 98 10 1 9.3×10−3 1.0×10−3 1.1×10−4 table 4. probability p(vn > vd) of exceedance the design values vd. figure 3. variation of the target probabilities ptar,t with reference time t in years for basic probabilities for one year ptar,1 = 10−5, 10−6 and 10−7. 7. concluding remarks (1) a tornado may be for any structure very dangerous depending on the extent of the structure and tornado and its intensity. tornado is unpredictable and of random time behaviour. (2) the probability of a tornado strike can be assessed taking into account the extent of a structure and tornado and the extent of the reference territory where tornados are observed. (3) tornado intensity expressed by wind speed is determined after the tornado strike following observed consequences. (4) probability structural failure is given by a product of the probability of a tornado strike and the probability of wind speed exceedance of design value. (5) estimation of wind speed exceedance of design value is strongly dependent on assumed wind speed distribution and the number of tornado occurrences during the lifetime of the structure. (6) available documents provide for target probability for structural failure value 10-7 per year, however, it is unclear how this criterion has been derived. (7) further reliability analysis of structures stricken by tornados is to be focused on detailed analysis of tornado consequences, on wind speed distribution and the number of tornado occurrences during the lifetime of the structures. references [1] iaea safety guides 50-sg-s11a. extreme meteorological events in nuclear power plant siting excluding tropical cyclones. international atomic energy agency, vienna, austria, 1981. [2] pacific northwest national laboratory. tornado climatology of the contiguous united states. nureg/cr-4461, rev. 2, u.s. nuclear regulatory commission office of nuclear regulatory research, washington, 2007. [3] h. zhu, j. chen, f. li, et al. tornado hazard assessment for a nuclear power plant in china. energy procedia 127:20-8, 2017. https://doi.org/10.1016/j.egypro.2017.08.091. [4] m. holický. reliability analysis for structural design. sun press, stellenbosch (south africa), 2009. [5] m. holický. introduction to probability and statistics for engineers. springer, heidelberg, 2013. 5 https://doi.org/10.1016/j.egypro.2017.08.091 acta polytechnica ctu proceedings 36:1–5, 2022 1 introduction 2 probability of structural failure 3 design wind speed 4 number of tornado strikes 5 failure probability 6 target probability of failure 7 concluding remarks references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.35.0042 acta polytechnica ctu proceedings 35:42–48, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague comparison of selected parameters for evaluation of rail surface damage intensity jiří šlapák∗, tomáš michálek university of pardubice, faculty of transport engineering, department of transport means and dignostics, studentská 95, 532 10 pardubice, czech republic ∗ corresponding author: jiri.slapak@upce.cz abstract. this paper deals with the issue of evaluation of a rail surface damage (rsd) intensity. some ways of calculating parameters that represent the rsd are described. in this context, a multi-body model of a railway vehicle was created and several simulations of this model on a curved track were performed. furthermore, these simulations were evaluated and the rsd parameters were compared. keywords: wheel/rail interaction, rail surface damage, wear number, multi-body simulation. 1. introduction in recent years, some railway infrastructure managers have started to use vehicle ratings based on damaging effects of vehicles on tracks. these damaging effects are directly related to maintenance requirements of tracks. one of the damaging effect is the rail surface damage (rsd), which occurs when wheels roll on rails. rsd primarily represents a wear of rails by abrasion and secondarily a relationship between the wear and a rolling contact fatigue (rcf ). this method of evaluation may motivate vehicle operators to use and purchase track-friendly vehicles. this paper is focused on several methods for evaluating rsd intensity and these methods are compared. these evaluation methods are described. the parameters that represent rsd intensity are compared depending on selected vehicle parameters. by means of multi-body simulations of vehicle running, a quantification of the parameters representing rsd intensity and damaging effects can be performed. 2. rail surface damage rolling of wheels on rails is possible due to normal forces acting in wheel/rail contact areas and the existence of adhesion in these contact areas. in general, when wheels rolls on rails, creepages and tangential (creepage) forces occur in the wheel/rail contact areas. the creepages occur in the longitudinal and lateral directions and also as a spin (rotation around the vertical axis). the longitudinal creepage is primarily related to traction and braking forces and also to the conditions of the wheelset/track interaction (specifically on the delta-r function) in curves. the increase in the lateral creepage is caused by an increase in the angle of attack of the wheelset. the spin is related to the inclination of the wheel/rail contact area. creepages and spin result in tangential (creepage) forces and spin moment that cause loading of the rail surface. due to this loading, the rail surfaces are damaged. in more detail, the issue of creepage forces is described in [1]. the first damaging effect is the wear of the rails and wheels by abrasion. the amount of wear depends on some design parameters of a vehicle, some wheel/rail contact conditions (coefficient of friction/adhesion, materials) and some track parameters. the rolling contact fatigue (rcf ) is another damaging effect caused by wheels rolling on rails. rcf causes cracks on the rail surface. under certain conditions, crack initiations can be removed by wear on the rail surface. thus, the wear can be beneficial in terms of the rolling contact fatigue. 2.1. wear number the parameter called the wear number tγ [nm/m] is the first option for evaluating and comparing the damaging effects of vehicles that result in rsd. this parameter is based on the physical assumption that the wear of the rails and wheels is caused by friction work performed in the wheel/rail contact. according this assumption, the wear number is defined as: tγ = |txγx| + |ty γy | , (1) where t [n] is the tangential (creepage) force and γ [−] is the creepage in the wheel/rail contact. the letters x and y describe the longitudinal and lateral direction of these quantities. equation 1 applies when spin and spin moment are neglected. since these quantities (specifically creepages and lateral creepage force) cannot be measured on a real vehicle, it is necessary to determine the value of the wear number using multi-body simulations of vehicle running. the wear number presented in equation 1 corresponds to the specific friction work performed in the wheel/rail contact. furthermore, the wear number in this form is used in the methodologies of some railway infrastructure managers for setting track acess charges (e.g. methodology [2]). 42 https://doi.org/10.14311/app.2022.35.0042 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 35/2022 comparison of selected parameters 2.1.1. rcf prediction method based on tγ the wear number tγ is also used in the rcf prediction method. the non-linear dependence of the wear number and the so-called rcf damage index is shown in figure 1 and described in [3]. this index indicates whether the rails are damaged due to wear, rcf or a combination of both, which is more common option. according to figure 1: • rcf damage index increases from 0 to 1 · 10−5 as the wear number values incease from 15 n to 65 n. when the wear number value is 65 n, the probability of rcf crack initiation is greatest. • then the index value decreases to 0 as the wear number increases to 175 n. in this part, a wear begins to predominate over rcf. • for the wear number values greater than 175 n, the rcf damage index has negative value. this mean that only wear damage occurs. figure 1. dependence of the wear number tγ and rcf damage index. [3] figure 1 applies to r260 steel. for other steels, the position of the characteristic points differs. 2.2. rail surface damage parameter according to en 14363 in the standard en 14363 [4], the current evaluation of rail load in lateral direction is performed using the quatistatic lateral guiding force ya,qst. this force is also used as indirect evaluation parameter for the rail surface damage intensity especially the wear of rails, but sometimes it shows a very weak connection with rsd. another parameter for the evaluating of the rail surface damage intensity is presented in annex k of standard en 14363 [4]. the standard proposes the parameter tqst which is a combined quantity of lateral yqst, longitudinal tx,qst and vertical qqst forces acting in the wheel/rail contact and represents the rail surface damage intensity. the parameter tqst is defined as: tqst = qqst 10000 · ( 330 · f 2 − 62 · f + 4 ) , (2) where f = yqst qqst + 0, 62 · |tx,qst| qqst . (3) the constants in these equations are derived as regression parameters from the dependence of tqst and tγ . the parameter f (equation 3) is dimensionless then the unit of the parameter tqst (equation 2) is newton [n]. this parameter has been defined in order to be able to measure the values of input parameters (forces) on a real vehicle without multi-body simulation. compared to the guiding force ya,qst, the parameter tqst better includes the influence of friction conditions in the wheel/rail contact area. the parameter tqst has only been defined for the guiding wheel of a vehicle. 3. multi-body simulations the values of previously defined quantities were obtained from multi-body simulations of running vehicle. the model and multi-body simulations of railway vehicle running were performed using simpack simulation software. 3.1. model of railway vehicle for the purposes of this study, the multi-body model of a conventional passenger car of an electric unit was used. the fundamental parameters of this model are listed in table 1. parameter value unit nominal mass of the carbody 40000 kg mass of the bogie frame 5200 kg mass of the wheelset 2100 kg bogie distance 19 m nominal bogie wheelbase 2.4 m vertical stiffness of the primary spring (per wheel) 1.6 kn/mm vertical stiffness of the secondary spring (per side of bogie) 0.7 kn/mm table 1. fundamental parameters of the multi-body model of the railway vehicle. setting the wheel/rail contact conditions is a very important part of the simulations. the wheel profile ore s1002 and the rail profile 60e1 were used. the rail inclination of 1:40 was considered. the nominal friction coefficient value 0.4 was chosen. the fastsim algorithm was chosen to calculate the tangential (creepage) forces. figure 2. the vehicle model created in simpack. 43 jiří šlapák, tomáš michálek acta polytechnica ctu proceedings figure 3. basic comparison of the mentioned parameters used to evaluate rsd intensity depending on the curve radius r. only for the guiding wheel on vehicle. 3.2. track conditions the simulations were performed on several curved tracks with a radius from 250 m to 1200 m. the length of the curve was set to 600 m. the cant d and cant deficiency i values are constant for all simulations. • d = 150 mm • i = 130 mm according to these conditions, the vehicle speed was set in the range of values from 77 km/h to 169 km/h. in the order to consider a real model of the track, reference track irregularities were used and a model of elastic track foundation was created. 4. simulation results time records of the quantities acting in the wheel/rail contact (lateral, longitudinal and vertical forces) and the parameter wear number tγ were monitored and exported from the multi-body simulations. from these records, the mean values of the quantities in the fully curved part of the tracks were calculated. the exported data from simulations were processed in matlab. because the parameters tqst and ya,qst are defined for the guiding wheel of a vehicle, all comparisons and evaluations have been processed for the guiding wheel only. 4.1. comparison of evaluation parameters for the first look at the comparison of the mentioned rsd intensity parameters (tγ , tqst, ya,qst), the parameter values depending on the curve radius are plotted in figure 3. this figure only shows the situation on the guiding wheel of vehicle for the nominal setting of the multi-body model and the simulation. according to figure 3, all rsd intensity parameters depending on the curve radius have the same trend and the shapes of the graph curves are similar. the values of the parameters progressively increase as the curve radius decreases. the figure further shows that the parameter tqst values are smaller than the wear number tγ values in curve radii greater than 700 m. then for smaller values of the curve radii, the parameter tqst values are greater then the values of the wear number tγ . however, this only applies to the guiding wheel of the vehicle. figure 3 further shows that the curve of the wear number tγ has a convex character in the whole range of curve radii. this also applies to the parameter tqst, except for the situation of the very small curve radius where the curve begins to be concave. 4.2. influence of selected parameters according to the standard en 14363 [4], the formula for the parameter tqst (equation 2) was defined for a wide range of selected operating conditons of a vehicle. therefore, vehicle running simulations were performed with a focus on the influence of these parameters: • friction coefficient in the wheel/rail contact µ, • weight of vehicle body m, • bogie wheelbase 2a, • longitudinal stiffness of wheelset guiding kx per one wheel. for all these parameters, the dependences of tqst and tγ on the curve radius r are plotted in graphs. to compare of rsd intensity parameters, the difference between tqst and tγ is defined and calculated as: tγ − tqst [n ]. (4) furthermore, the relative difference is defined and calculated as: (tγ − tqst) /tγ · 100 [%]. (5) 4.2.1. influence of friction coefficient in the first part, the influence of the friction coefficient values in the wheel/rail contact was investigated. friction coefficient values of 0.2, 0.4 and 0.6 were selected. figure 4 shows comparison of the wear number tγ and the parameter tqst (calculated according to [4]) on the guiding wheel of the vehicle. for different 44 vol. 35/2022 comparison of selected parameters values of the friction coefficient µ in the wheel/rail contact area, the individual curves are plotted as a function of the curve radius r. for all investigated values of the friction coefficient, this figure shows that the values of the parameter tqst are very close in the range of the curve radius from 700 m to 1200 m. this behavior of the parametr tqst is different from the wear number tγ . further, as the curve radius value decreases, the values of the parameter tqst as well as the parameter tγ values increase. figure 4. dependence of the tqst and tγ on the curve radius r in defined range of the friction coefficient values µ. figure 4 shows that the friction coefficient value in the wheel/rail contact area has a great influence on the values of wear number tγ and also shows that the parameter tqst follows this trend. figure 5. dependence of the ya,qst on the curve radius r in defined range of the friction coefficient values µ. in figure 5, the influence of the friction coefficient µ on lateral guiding force ya,qst is shown. as the friction coefficient values increase, the slope of the plotted curves increases. this behavior is the same as for the parameters in figure 4. but the curves are shifted, which causes a large mismatch between the force ya,qst and the parameters tqst and tγ . the next figure 6 shows the difference and the relative difference between the values of the wear number tγ and the parameter tqst in defined range of friction coefficient values. in general, the best match of these parameters occurs for the smallest value of the friction coefficient. but for very small curve radii, the value of the relative difference for the friction coefficient of 0.6 is the smallest. this mean that in very small curve radii, tqst corresponds better to the wear number tγ with increasing value of the friction coefficient. on the other hand, for large curve radii, the parameter tqst best corresponds to the wear number tγ under the conditions of the smallest value of the friction coefficient. for a radius curve of 1200 m and a friction coefficient of 0.6, the value of the relative difference is 78%. figure 6. the difference tγ − tqst depending on the curve radius r (above) and the relative difference (tγ − tqst) /tγ depending on the curve radius r (below) in the defined range of friction coefficient µ values. 4.2.2. influence of vehicle weight in the next part, the influence of the vehicle weight was investigated. for this evaluation, the vehicle body mass m was set to 35, 40 and 45 tons. figure 7 shows the values of the parameter tqst and the wear number tγ as a function of the curve radius r for the defined values of the vehicle body weight. according to this figure, the weight of vehicle has an effect on the wear number tγ values. this effect is most evident for the small curve radii. on the other hand, the values of the parameter tqst do not change when the vehicle weight changes. this applies in whole range of the curve radii expect the very small curve radii. 45 jiří šlapák, tomáš michálek acta polytechnica ctu proceedings figure 7. dependence of the tqst and tγ on the curve radius r for defined values of the vehicle body mass m. figure 8. dependence of the ya,qst on the curve radius r for defined values of the vehicle body mass m. figure 8 shows that the weight of the vehicle body has an effect on the guiding force ya,qst. the wear number tγ (in figure 7) has a similar effect. furthermore, the dependence of the force ya,qst on the curve radius is concave in very small curve radii, as well as the dependence of the parameter tqst on the curve radius. in figure 9, the difference and the relative difference of the parameters tqst and tγ are shown in relation to figure 7. in very small curve radii, the best match of the investigated parameters occurs at higher values of the vehicle weight. then in the large curve radii, the best match of the parameters occurs for the lightest vehicle but the differences disappear in very large curve radii. in the value of the curve radius of 1200 m, the relative difference value is 72% and this value does not depend on the vehicle weight. figure 9. the difference tγ − tqst depending on the curve radius r (above) and the relative difference (tγ − tqst) /tγ depending on the curve radius r (below) for defined values of the vehicle body mass m. 4.2.3. influence of bogie wheelbase in this part the influence of the bogie wheelbase 2a on the parameters tqst and tγ is analyzed. the value of the bogie wheelbase was set to 2.0, 2.4 and 2.8 meters. figure 10. dependence of the tqst and tγ on the curve radius r for defined values of the bogie wheelbase 2a. in figure 10, the influence of the bogie wheelbase values is evident for both parameters of rsd intensity over the whole range of the curve radii. again, the values of tqst are smaller than the values of tγ for large curve radii. then, for small curve radii, the values of tqst increase over the values of tγ . 46 vol. 35/2022 comparison of selected parameters figure 11. dependence of the ya,qst on the curve radius r for defined values of the bogie wheelbase 2a. according to figure 11, the influence of the bogie wheelbase 2a on the guiding force ya,qst values and on the parameter tqst is similar. for large and very small curve radii, it is difficult to predict and describe the values of the force ya,qst . the dependence of the tγ values on the curve radius has the same trend for different values of bogie wheelbase. this does not apply to the parameter tqst whose trends are changing in the very small curve radii. this is also shown in figure 12 where the both differences increase (in terms of absolute values) in the very small curve radii under the condition of the bogie wheelbase of 2.0 m. but for the other values of the bogie wheelbase in the very small curve radii, the differences decrease with decreasing the curve radius. figure 12. the difference tγ − tqst depending on the curve radius r (above) and the relative difference (tγ − tqst) /tγ depending on the curve radius r (below) for defined values of the bogie wheelbase 2a. in general, figure 12 shows that the best match between the tqst and tγ parameters occurs for the vehicle with the longest bogie wheelbase. according to the first graph in figure 12 and for the curve radius of 250 m, the value of the difference is 150 n for the vehicle with bogie wheelbase of 2.0 m. it is the highest value of the absolute diference of the rsd intensity parameters. but from the point of view of the relative difference of these parameters, the value corresponds to the relative difference of 37%. a much larger relative difference occurs in curve radius of 950 m where the value is about 74% and it is the highest value of the relative difference of the parameters for this analysis. 4.2.4. influence of longitudinal stiffness of wheelset guiding the influence of the longitudinal stiffness of the wheelset guiding kx per one wheel on the parameter tqst and the wear number tγ is the last analyzed part. the values of the longitudinal stiffness of the wheelset guiding per one wheel were set to 2.0 · 107, 3.5 · 107 and 5.0 · 107 n/m. figure 13. depencence of the tqst and tγ on the curve radius r in defined range of the longitudinal stiffness of wheelset guiding kx per one wheel. figure 14. depencence of the ya,qst on the curve radius r in defined range of the longitudinal stiffness of wheelset guiding kx per one wheel. figure 13 shows that the longitudinal stiffness of the wheelset guiding in defined range of values has 47 jiří šlapák, tomáš michálek acta polytechnica ctu proceedings a small effect (in comparison with other investigated parameters) on the values of the parameter tqst and the wear number tγ . from this point of view, the behavior of these parameters is the same. thus, in general, changing the stiffness value causes the same reaction in the parameter tqst as in the wear number tγ . according to figure 14, the behavior of the dependence of the guiding force ya,qst on the curve radius is the same as the dependences on figure 13. this means that the effect of the longitudinal stiffness of wheelset guiding kx in a defined range of values on the guiding force ya,qst values is small. figure 15. the difference tγ − tqst depending on the curve radius r (above) and the relative difference (tγ − tqst) /tγ depending on the curve radius r (below) in the defined range of the longitudinal stiffness of wheelset guiding kx per one wheel. this fact is confirmed in figure 15. this figure shows that the difference and the relative difference values are almost idential for the defined range of the longitudinal stiffness values. some effect of the longitudinal stiffness to the difference and the relative differrence of rsd intensity parameters occurs in very small curve radii. 5. conclusions it can be assumed that the wear number tγ represents the wheel and rail abrasion wear and indirectly also represents the rolling contact fatigue effects. it is used by railway infrastructure managers as an indicator of damage and maintenance requirements of a curved track. unfortunately, the wear number must be obtained from multi-body simulations and cannot be measured on a real vehicle. this fact means a risk that the results of the multi-body simulations can be affected by the model settings. the correctness of this settings cannot be experimentally verified. the lateral guiding force ya,qst is a parameter that is used to evaluate the lateral load of the rails. this force exhibits similar behavior and trends as the wear number tγ depending on the curve radius, but its values are different. for the reason, an alternative quantity, the parameter tqst is defined in the standard en 14363. based on multi-body simulations of running a conventional passenger four-axle railway vehicle, it was found that the parameter tqst shows the same trends as the wear number tγ on the outer guiding wheel of a vehicle. however, there are differences between the values of these parameters, which vary depending on the curve radius. the parameter tqst values are smaller than the wear number values in the large curve radii. conversely, the parameter tqst values are higher than the wear number in the very small curve radii. the best match of these parameters occurs in curve radii from 400 to 600 meters. the worst match occurs in the very large curve radii where the relative difference between the parameter tqst and the wear number tγ has a value higher than 70 percent. this mean that this approximation of the parameter tqst is bad for very large curve radii. the friction coefficient value in the wheel/rail contact area has the greates influence on the rsd intensity parameters. further, the vehicle body weight affects the wear number values, but has almost no effect on the parameter tqst value. the values of the longitudinal stiffness of wheelset guiding have the same effect on both rsd intensity parameter. the defined parameter tqst for evaluation of rsd intensity follows the trend of the wear number. but between these parameters, there is bad match in the large curve radii and the parametr tqst does not respond to the vehicle weight change. the parameter tqst can represent the wear number, but the accuracy of this parameter could be improved. acknowledgements this work was supported by the scientific research project of the university of pardubice no. sgs_2021_010. references [1] voltr, p.: calculation of locomotive traction force in transient rolling contact. in: applied and computational mechanics, 11, 69-80, 2017. [2] feredal office of transport. base price wear in the train-path pricing system 2017 – instruction for determining vehicle prices. bern, 2017. [3] iwnicki, d. s.: the effect of profiles on wheel and rail damage. in: international journal of vehicle structures & systems,1(4), 99-104, 2009. [4] en 14363:2016+a1:2018. railway applications – testing and simulation for the acceptance of running characteristics of railway vehicles – running behaviour and stationary tests. brusel: european committee for standardization, 2018. 48 acta polytechnica ctu proceedings 35:42–48, 2022 1 introduction 2 rail surface damage 2.1 wear number 2.1.1 rcf prediction method based on t 2.2 rail surface damage parameter according to en 14363 3 multi-body simulations 3.1 model of railway vehicle 3.2 track conditions 4 simulation results 4.1 comparison of evaluation parameters 4.2 influence of selected parameters 4.2.1 influence of friction coefficient 4.2.2 influence of vehicle weight 4.2.3 influence of bogie wheelbase 4.2.4 influence of longitudinal stiffness of wheelset guiding 5 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0417 acta polytechnica ctu proceedings 38:417–423, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague characteristics of business models for innovation clusters in decarbonation projects matthias haasea, ∗, emilia motoascab a zurich university of applied sciences, institute of facility management, grüental / ra, p.o. box, ch-8820 wädenswil, switzerland b vito nv, boeretang 200, 2400 mol, belgium ∗ corresponding author: matthias.haase@zhaw.ch abstract. roughly 97 % of the european union (eu)’s building stock is not considered energy efficient, and 75 to 85 % of it will still be in use in 2050. residential buildings account for around two thirds of final energy consumption in european buildings. the rate at which new buildings either replace the old stock, or expand the total stock, is about 1 % per year. similarly, the current renovation rate of existing buildings in the eu is about 1–2 % of the building stock renovated each year. the transformation of today’s electric power sector to a more sustainable energy production based on renewable energies will change the structure of the industry. in this transformation towards a smart energy system interaction between sectors and technologies the main stakeholders (energy service providers; utilities) will face new challenges in their traditional way of doing business. therefore, adapting their business models to remain competitive is seen as an important step. we chose to characterize these business models by content, structure and governance of transactions for creating value by exploiting business opportunities. in the energy sector the following characteristics to the business models for energy supply have been identified. the most predominant archetypes of business models for the energy supply are presented and discussed. further, we propose to set up (or use existing) innovation clusters, based on these promising bm to ensure that innovative business environments (innovation clusters) will grow that have the potential for upscaling and replication of district decarbonization solutions. there are no specific business models for energy supply applied to renovation of districts. uncertainties in the supportive measures for the application of der makes it difficult to develop new business models for the utilities. keywords: business models, energy supply, innovation cluster. 1. introduction renovation strategies on building level need to be derived as a combination of energy efficiency upgrades for buildings and the use of renewable energy to decarbonise the energy supply, on district or city scale. to this end, the annex 75 of the iea ebc programme sets off to define a methodology to identify which strategies are most energy saving and cost-effective when applying both energy efficiency and renewable energy measures [1]. by combining energy efficiency and renewable energy sources, the approach addresses both energy supply and demand in the built environment. in this sense, building retrofitting is an appropriate strategy to reduce demand, while the use of renewable energy aims at decarbonizing the energy supply system. nevertheless, to apply the large-scale renovation strategies and achieve the projected building stock decarbonisation, identifying the technical solutions is not enough. the renovation rate in europe remains well below the targeted annual 3 % [2, 3]. some of the main barriers to renovation have to do with the renovation cost and access to finance, as well as complexity, awareness, stakeholders’ management and fragmentation of the supply chain [4]. as a result, business models are relevant to implementation and acceleration of renovations. seddon et al. [5] defines “business model” as the outline of essential details of a firm’s value proposition for its various stakeholders, and the activity system the firm uses to create and deliver this value proposition. in other words, a business model is the abstraction of a strategy, focused on the system of activities through which a firm creates economic value. 1.1. energy supply grid renovation strategies on district level comprise of energy efficiency measures on building level as well as energy supply measures on the district level. energy supply for the buildings stands in for the supply of both electric and thermal energy. most of the energy use of buildings is related to space heating (and cooling), therefore the needed thermal energy can be extracted directly from a dh(c)-network or provided by a heat pump (hp) or even directly converted using electric heaters/coolers. energy supply companies are responsible for supplying buildings and district with their needs in terms of electric and thermal energy, however still most of the provided energy 417 https://doi.org/10.14311/app.2022.38.0417 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en matthias haase, emilia motoasca acta polytechnica ctu proceedings figure 1. illustration of energy supply system options (blue: gas, green: renewable: dotted: new connections). is mostly electric energy. electricity supply companies today face a strong decrease in the number of customers. this has to do with new market actors which are operating with new business models centered on renewable energy and energy efficiency technologies. to understand the situation better, it is helpful to acknowledge that electricity supply companies act in an organized electricity market based on different marketplaces. the physical electric energy system encloses the infrastructure (generation, transport, distribution and use together with their components). thus, the market consists mainly of the following actors [6]: • electricity generator: who generates electricity and sell it to the energy suppliers. • electricity suppliers who purchase the electricity from the generators and sell it to consumers. • consumers who use electricity and pay monthly fees to suppliers. • transmission system operators (tso), who are responsible for transporting electricity for long distance and ensuring grid stability and reliability by real time dispatch. • distribution network operators (dso), who are responsible for delivering electricity to the consumers and measuring the consumption. • regulators, who set the market rules and oversee the functioning of the market. 1.2. energy supply value stream in the eu countries some energy supply companies operate transnationally, however still a large part of the worlds’ electric energy supply is either based on fossil fuels like coal, gas, and oil or nuclear energy. the production, transmission and distribution of electricity accounts for the largest share of the world’s anthropogenic greenhouse-gas emissions, while the use of emission-free nuclear energy comprises serious security risks and unsolved problems of hazardous waste, therefore the role of renewable energies as the most important instrument to mitigate climate change and reduce negative effects of energy production is increasing. despite the fact that utilities (with national or transnational activities) still have a dominant position they are confronted with disruptions of their current way of doing business and face the challenge to develop new business models for electricity generation from distributed and highly intermittent renewable sources. in this sense, the energy supply company changes into a service provider based on the idea of an innovation cluster which can be understood in two ways. the traditional view is that it is a collection of companies situated within some level of proximity, allowing for more collaboration, interaction, development of stronger ties and a natural growth of a collaborative strengths within the cluster. in our work, these new energy service providers or innovation clusters acts as ecosystems that create an active flow of information and resources for ideas to transform into reality. through these ecosystems, a process is started which creates expertise in new areas (here buildings and energy), helps to diversify the economy, and allows businesses to meet their customers where they are. additionally, an innovation ecosystem provides the means to create economic stability and resource sharing [7]. the value of an innovation ecosystem lies in the access to resources and the flow of information for the ecosystem’s stakeholders. this information flow 418 vol. 38/2022 characteristics of business models for innovation clusters . . . exists already in some energy supply companies and creates more investment opportunities for the right institutions to connect with the right ideas for their businesses and portfolios, at the right time, for the right reasons. for energy supply companies which want to activate value streams in the building sector it is important to understand the different business models. 1.3. research question the transformation of today’s electric power sector to a more sustainable energy production based on renewable energies will change the structure of the industry [8]. in this transformation towards a smart energy system interaction between sectors and technologies the main stakeholders (energy service providers; utilities) will face new challenges in their traditional way of doing business. therefore, adapting their business models to remain competitive is seen as an important step. when looking at the business model literature reveals that there exist basically two possibilities: (1.) ownership of renewable energy assets [9]. (2.) utilities need to develop from commodity providers to energy service providers [10, 11]. according to this idea utilities should evolve to comprehensive energy-solutions providers for residential and commercial customers to create new sources of revenues. but which characteristics of these new value creation streams exist and which could be extended and developed further? 2. methodology new business models are needed that are able to work with energy demand reduction and supply. literature review was conducted in the scopus and science direct database based on the method iterative method proposed by bocken et al. [12] starting with a cascading keyword search (business model, energy supply, districts, characteristics) resulting in 844 publications. of these were 69 review articles of which 50 were reviewed to identify prominent schemes. of the 710 research articles, 241 were doublets and disregarded resulting in a list of 468 research articles. a total of 523 articles were reviewed to identify examples and case studies. a total of 34 in-depth interviews with different stakeholders were conducted between november 2020 and september 2021 with energy experts from the netherlands, switzerland, span, austria and germany. this formed the basis for compiling a number of main characteristics of business models for energy supply. we chose to characterize these business models by content, structure and governance of transactions for creating value by exploiting business opportunities. “content” refers to the performed activities (transactions/exchanges of goods and information), together with the required capabilities and resources. “structure” describes the links among the activities (transactions) and the participant parties, while “governance” provides information about who performs the activities and where, thus it refers to legal organizational form, incentivizing of participants and the way information, goods and resources flow. the analysis and compilation of the most common business models for energy supply can help to develop their own bm following the bmc (business models canvas) principle. 3. results in the energy sector the following characteristics to the business models for energy supply have been identified that can be characterized by servitization (energyas-a-service), financing and ownership, public interest, customer’s role, decentralization, energy tariff’s structure, infrastructure, flexibility, and management and control. 3.1. servitization (energy-as-a-service) the bm have a more service-oriented character, thus the companies are selling the functionality of the product rather than the product energy. the product is thus replaced by a combination of products and services and the notion of the value moves from exchange to application. in the energy transition context, servitization is correlated with energy services and energy efficiency. this means the perception of having a certain percentage cost savings due to the reduction of end-user’s energy consumption. the variations of energy services range from additional services such as aggregated information and provision of data analysis to advanced services, such as performance and activity control, depending on servitization intensity. energyas-a-service activities include energy planning, project design, implementation, maintenance, management, operation evaluation and equipment supply. in this bm, performance is directly linked to the energy savings achieved where its remuneration includes savings guarantees. 3.2. financing and ownership for the energy supply (mainly of the renewable energy sources (res)) three main ownership models are predominant: consumer’s ownership, collective community ownership and service-based with company ownership [13]. since res capacity may have an influence on the grid stability and energy supply security, the value stream is based on the notion that owners may influence grid balance. one emerging bm is the remuneration of these services to increasing or decreasing grid stability. in some bm there is a clear distinction between the customer or/and third-party ownership, where the prosumers or aggregators can choose to contribute (or not) to grid balance activities. this relates to the utility ownership, i.e. the energy supply company has the full authority to manage and maintain the grid and control the renewables production. 419 matthias haase, emilia motoasca acta polytechnica ctu proceedings collective community ownership is often considered as a bm where the source of income can be controlled locally. therefore, financing these kinds of investments are more likely to be accepted within the community, because it helps developing local res ownership and avoids that value leaks out of the local economy. financing of res technologies is a crucial factor for both micro-generation (where the costs are barrier because of the long-term investment in the infrastructure assets and the success or failure of investment depends on the institutional support) or for large-scale res technologies (where the large upfront cost is often described as a “barrier that prevents customers from financing res and hence outsourcing financing to a third-party in order to remove this barrier”) [14]. alternative financing sources for re investments are the newly emerging energy cooperatives, where the financial risk can be mitigated by local community investment or collectively fundraising for re through crowdfunding platforms. 3.3. public interest the public sector may wish to steer a district energy project towards a variety of specific local objectives [15], including: cheaper local energy for public, private and/or residential customers (e.g., the alleviation of fuel poverty); local job creation; local wealth retention; low-carbon power generation; and/or local air pollution reduction. economic modelling helps quantifying these objectives and thus creating additional value by extending standard financial modelling. 3.4. customer’s role understanding the different roles of the customer is centrally important to reduce the cost in the energy supply and specifically in the demand side management (dsm) for compensating consumers for their participation in demand response programs [14]. the relationship with customers has been modified since some years, by intensifying customer engagement or by delivering new services, by providing consumption and production related information real-time or by installing enhanced two-way communication channels. the knowledge about consumer’s behavior, including attitudes, tastes and needs are critical success factors for the bm operating in decentralized systems. the reason is that multiple roles for the consumers are possible. first, there are active producers and consumers who produce and self-consume green electricity and/or heat simultaneously and alternating (e.g. day/night). thus, customers act as financial investors in res and demand user services as light, heat, etc. instead of an energy commodity. this results in interactions with local beneficiaries, project supporters/protestors/activists, and not least with technology hosts. 3.5. decentralization energy systems can be designed on different levels of decentralization. nowadays, due to the distributed nature of res and their smaller production capacity a new decentralized energy market has been established. it requires different revenue models than the classical ones of centralized energy production. the associated bms should provide distributed solutions (e.g. for each consumer separately), thus implying higher costs compared to previously dominating one-size-fitsall solutions. this provokes a strategy shift from big to small, from wholesales to a customer-oriented strategy, from commodity to service, and from long-term energy planning to a more flexible energy planning including further services/infrastructure like e.g. energy storage. in res dominated energy production, the ownership of the res and the vicinity between production and consumption sites plays a determining role in the choice of bm. 3.6. energy tariffs structure today, energy company’s revenues are typically based on a flexible fee and fixed price per energy. this means that the more energy is consumed the better it is for the utility. this is the reason why the current revenue model creates a incumbrance for utilities to offer incentives for energy efficiency, because thereby energy delivery decreases and revenues reduces. it is often argued that the current revenue model is the greatest obstacle between the current utility structure and a modernized energy delivery system or thirdparty owned decentralized power generation based on renewable sources [16]. here, a decoupling of the relationship between sales volume and revenues is needed, a dynamic pricing (see flexibility below) should orientate on the wholesale prices for energy (time-of-use tariffs, off/peak prices, etc.), flatrate tariffs should be coupled to incentives to reduce energy consumption, new const structures would need to be more customer oriented. 3.7. infrastructure the infrastructure includes key resources to create a value proposition. especially renewable energy assets follow different strategies than traditional assets (high investment, low operational costs) and depend on key activities of the energy provider (capacity), key partnership within a network of suppliers and partners to make the bm work. even co-operations or joint ventures between utilities and independent developers are needed to build up portfolios and inhouse competencies to extend the revenue model. in this respect, utility-side renewable energy business models are more attractive to utilities in terms of risk and return expectations than customer-side renewable energy business models. thus, when considering risk return expectations and transactions costs, energy companies will primarily focus on large scale projects. 420 vol. 38/2022 characteristics of business models for innovation clusters . . . archetype type novelty oriented energy bms going green bm building energy communities bm lock-in oriented business models bm that offer energy functionalities complementarities-oriented energy supply business models optimizing grid operations combining value propositions acting locally efficiency-oriented energy business models scaling-up running platforms table 1. summary of business models. hence, customer-side renewable energy business models are not expected to advance in scale in the near future [8]. 3.8. flexibility flexibility is the ability of energy supply systems to manage net load variation (and generation outage) with their existent resources. this spans over various time horizons when intermittent res, such as solar and wind energy are present in the system. flexibility can be stimulated either from consumption or from generation side by coupling them with timing service. the decentralized electricity and/or heat generation is thus not just developing sources of renewable energy but also finding ways of locally based balancing of production and consumption. trading of flexibility services are important for reliable energy supply systems, due to their main functions (integration of intermittent resources, congestion management and portfolio optimization) which affect different users. 3.9. management and control mc indicates the actors that take the responsibility of maintaining and keeping the energy supply hardware in optimal operating conditions. many factors affect management (consisting of operation, control and governance), among other the proximity of the technology to the consumption’s site, the contract, the partnership and the legal form. the management activities are actually optimization activities that improve electricity grid balance and trading service, maintain the co-owned infrastructure and lever the fluctuation of renewable energy production and grid balance. table 1 summarizes the types of business models that these innovation clusters can follow. there were basically four different business model archetypes identified which can be split in several types and even sub-types. however, there are other objectives that are to be fulfilled as well. buildings are more commonly seen as micro-energy hubs with energy generated, stored, used and saved in buildings and districts as indicated by bpie (2016), aiming at [17]: • maximising energy efficiency of the buildings. • increasing on-site or nearby res production and self-consumption. • encouraging energy storage capacities in buildings (or nearby). • incorporating demand-response capacity in the building. • decarbonizing the heating and cooling energy for buildings. • empowering end users via smart meters and controls. • making dynamic price signals available for all consumers. • fostering business models aggregating micro-energy hubs. • (re)building smart and interconnected districts (renovate/retrofit). • building up infrastructure for further market uptake of electric vehicles. the most predominant archetypes of business models for the energy supply are presented in table 1. • novelty-centered bms are the ones where new ways of performing the economic transactions have been adopted. novelty is rooted in one or several activity system design elements and present in different innovation forms. accounting for the content element, the fossil fuel energy is replaced in these bms with res, thus they are mostly technology driven bms. nowadays these bm provide a strong predominance in the solar pv businesses, resulting in a pattern category named “going green”. “building energy communities” is the second pattern category where new organizations based on co-participation are addressed in the structure element, while the governance element is based on shared resources and governance. • lock-in centered bms refer to the ability of the company to attract, maintain and improve customer and partner association with the bm. • complementarities-centered bms refer to the bm bundling goods and services together instead of providing each separately and finally 421 matthias haase, emilia motoasca acta polytechnica ctu proceedings • efficiency-centered bm are the ones where measures are taken to achieve increased transaction efficiencies. 4. conclusions the energy sector is undergoing a continuous process of transformation where, together with decentralization and digitalization, a ultimate shift of energy supply towards renewable, co2 neutral energies is taking place. the classical structure of the electrical energy industry that emerged after the liberalization of the electricity and gas markets in europe including established business models, is subject to disruptive and substantial changes. business models in the electric power sector have to be seen as interlinked with the regulatory and policy frameworks that characterize the sector. up to now, the energy system is based on a distribution system which stateor national government-appointed regulatory commissions regulate. this influences also the revenues of electricity distribution companies and thus the viability of distributed renewable energy businesses (der) in distribution networks are exposed in part to these regulatory frameworks. correspondingly, market rules are established by a central authority, independent system operators (isos) or regional transmission operators (rtos), all entities monitored and regulated by an energy regulatory commission (ferc) in wholesale electricity markets. in addition, the electric power sector is subject of significant national and eu policy support, taking the form of subsidies or favorable rules for a variety of technologies, like e.g. solar-based energy (electricity and heat) generation. on the other hand, new der business models selling services in wholesale electricity markets must conform to the market rules and regulations established by these authorities. one way out would be to change these market rules and regulations together with authorities and policy makers. better understanding these regulatory and policy interdependencies is decisive to ensure the sustainable development of these businesses. • four different archetypes of bm were identified that ensure a maximum of impact. • there are no specific business models for energy supply applied to renovation of districts. • uncertainties in the supportive measures for the application of der makes it difficult to develop new business models for the utilities. acknowledgements this work has been developed 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[2022-10-12]. https: //www.bpie.eu/publication/buildings-as-microenergy-hubs-delivering-climate-solutions/ 423 https://www.techstreet.com/ashrae/standards/or-20-020-energy-master-planning-identifying-framing-constraints-that-scope-your-technology-options?product_id=2113219 https://www.techstreet.com/ashrae/standards/or-20-020-energy-master-planning-identifying-framing-constraints-that-scope-your-technology-options?product_id=2113219 https://www.techstreet.com/ashrae/standards/or-20-020-energy-master-planning-identifying-framing-constraints-that-scope-your-technology-options?product_id=2113219 https://www.techstreet.com/ashrae/standards/or-20-020-energy-master-planning-identifying-framing-constraints-that-scope-your-technology-options?product_id=2113219 https://www.renewableenergyworld.com/baseload/renewable-energy-and-the-utility-the-next-20-years/#gref https://www.renewableenergyworld.com/baseload/renewable-energy-and-the-utility-the-next-20-years/#gref https://www.renewableenergyworld.com/baseload/renewable-energy-and-the-utility-the-next-20-years/#gref https://www.bpie.eu/publication/buildings-as-micro-energy-hubs-delivering-climate-solutions/ https://www.bpie.eu/publication/buildings-as-micro-energy-hubs-delivering-climate-solutions/ https://www.bpie.eu/publication/buildings-as-micro-energy-hubs-delivering-climate-solutions/ acta polytechnica ctu proceedings 38:417–423, 2022 1 introduction 1.1 energy supply grid 1.2 energy supply value stream 1.3 research question 2 methodology 3 results 3.1 servitization (energy-as-a-service) 3.2 financing and ownership 3.3 public interest 3.4 customer's role 3.5 decentralization 3.6 energy tariffs structure 3.7 infrastructure 3.8 flexibility 3.9 management and control 4 conclusions acknowledgements references my title https://doi.org/10.14311/app.2022.36.0191 acta polytechnica ctu proceedings 36:191–197, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague structural reliability of existing rc beams strengthened with uhpfrc tensile layers lenganji simwanda∗, nico de koker, celeste viljoen, adewumi john babafemi stellenbosch university, department of civil engineering, private bag x1, matieland 7602, south africa ∗ corresponding author: 24429902@sun.ac.za abstract. a methodology for reliability analysis of reinforced concrete (rc) beams strengthened with ultra high-performance fibre reinforced concrete (uhpfrc) tensile layers is presented. the proposed methodology includes stochastic stress-block analysis of a section, assuming a perfect bond between the rc beam and the uhpfrc layer. annual reliability analysis of the rc beam before and after the strengthening operation is conducted. deterioration induced by chloride corrosion is incorporated into the analysis via a chloride induced corrosion model based on fick’s law of diffusion and described stochastically to account for the epistemic uncertainty in the time to corrosion initiation and rate of corrosion. a plot for determining the required thickness of the uhpfrc tensile layer to upgrade to the required reliability level is also given, considering the time from construction to when the strengthening operation is conducted. the proposed approach is easy to apply for routine practice. keywords: corrosion damage, rehabilitation of existing structures, reliability analysis, uhpfrc tensile layers. 1. introduction addition of ultra high-performance fibre-reinforced concrete (uhpfrc) layers on the tensile side is a newly developed technique for strengthening reinforced concrete (rc) constructions, such as floor slabs, beams and bridge decks [1]. other conceptual ideas include additional layers on the compression side or even including on the sides (referred to as jackets) to improve the bending resistance of rc beams. efforts have been made in the recent past to investigate the efficacy of the uhpfrc strengthening techniques for retrofitting and rehabilitation of existing rc structures. for example, farhat et.al [2] observed an increase in the load carrying capacity and reduction in autogenous shrinkage strains of damaged rc beams retrofitted with uhpfrc and proposed a load-carrying prediction model based on fracture mechanics. habel et.al [3] investigated, through a parametric study, the flexural response of composite uhpfrcrc elements via the traditional analytical crosssectional approach and observed that addition of steel reinforcement in the uhpfrc layer is the most effective approach to increase the moment capacity. a most recent work by lampropoulos et.al [4] assessed, experimentally and numerically, the efficacy of uhpfrc layers and jackets for the strengthening of existing rc beams, and observed that the three side uhpfrc jacket approach was superior in terms of performance. with efforts to apply these techniques to routine practice, analytical and numerical models [1–4] have been proposed for the estimation of bending moment capacity of rc beams strengthened with uhpfrc layers. despite several uncertainties present in the formulations of the present analytical approaches, the application of structural reliability method has received very little attention in this research area. the objectives of this paper are two-fold. the first being to assess the viability of the stress-block section analysis in the estimation of nominal bending moment capacity of rc beams strengthened on the tensile side with uhpfrc layers, and the second being to present and apply a statistical methodology for conducting reliability analysis of existing rc beams strengthened with uhpfrc tensile layers. 2. stochastic modelling of deterioration of rc beams prior to strengthening 2.1. corrosion modelling a significant factor to the structural deterioration of rc beams exposed to saline conditions is the corrosion of steel reinforcements. therefore, their remaining service life when strengthened with uhpfrc is considered in this study in the context of chloride induced corrosion. the effect of corrosion-induced deterioration on the cross-section area of steel rebars is discussed while ignoring its effects on other geometrical and material properties of steel and concrete. for simplicity, a model that assumes uniform corrosion and proposed in [5, 6] is considered here, but more detailed corrosion models may be applied in appropriate cases following the methodology presented in this paper. in [5, 6], a model is presented such 191 https://doi.org/10.14311/app.2022.36.0191 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en l. simwanda, n. de koker, c. viljoen, aj. babafemi acta polytechnica ctu proceedings that the evolution in time of the diameter ϕ of the corroded rebars is computed via equation (1) ϕ(t) = { ϕinitial t ⩽ t0 ϕinitial − 0.0232icorr(t − t0) t > t0 (1) where ϕinitial is the initial bar diameter, t0 is the time (in years) to corrosion initiation, and icorr is the corrosion ratio (in µa/cm2) computed empirically via equation (2) icorr = 37.8(1 − w/c)−1.64 c′ (2) where w/c and c′ is the water to cement ratio of concrete and concrete cover to reinforcement, respectively. with this model, the reduction of the cross-section area of the steel reinforcement as(t) as a function of time can be computed from the time varying diameter ϕ(t). 2.2. corrosion stochastic simulation it is worth noting that equations (1) and (2) contain parameters associated with given levels of uncertainty, that needs to be quantified and accounted for in the reliability analysis of rc beams exposed to marine environments. specifically, the espistemic uncertainty in the time to corrosion initiation t0 and corrosion ratio icorr needs statistical description. following a previous work [7], the epistemic uncertainty in t0 and icorr were reported to be well described by weibull and uniform probability distributions, respectively. for the problem considered in this work, t0 was estimated to have weibull parameters (α = 11.3, βw = 4.81), whereas icorr was taken to be uniformly distributed between the closed interval icorr ∈ [2, 3]. figure 1 shows the mean time series and its corresponding 10% confidence interval of the stochastic process describing as and estimated via monte carlo sampling with 106 samples. 3. analysis of rc beams strengthened with uhpfrc 3.1. moment capacity: proposed approach the moment capacity of non-strengthened rc beams can readily be computed by analytical models reported in literature [8–10]. the eurocode model [9] of singly reinforced concrete beams is used in this work. building on the assumptions of the eurocode model [9], a procedure for estimating the moment capacity of rc beams strengthened with tensile uhpfrc layers is proposed. the approach is based on the stress-block analysis of a balanced section is, assuming a perfect bond between the rc beam and the uhpfrc layer and that the section cracks in the rc zone of the strengthened beam transferring the section forces to the longitudinal reinforcements and the fibres in the uhpfrc layers[11]. 10 20 30 40 50 0 50 100 150 200 250 300 figure 1. uncertainty quantification of deterioration of steel cross-sectional area in the rc beam over its lifetime. the analysis of the cracked section assumes strain compatibility across the cross-section (see figure 2), maximum strain in concrete, and yielding of steel fibres in the uhpfrc layer is reached first. with this assumption the characteristic moment capacity of the rc beam strengthened with layers may be estimated via the following equations. the stress σf and strain εf in the extreme steel fibres of the uhpfrc layer may be computed via equations (3) and (4) σf = 2τf lf df (3) εf = σf ef (4) so that the position of the the neutral axis of a balanced section can be obtained via a = fykas + ftkbtγ fckb (5) c = a/0.80. (6) the position where the force in the steel rebars and the force in the steel fibres effectively act (denoted as ζ) is ζ = fykas(h − c′ − c) + ftkbtγ(h − c − btγ) fykas + ftkbtγ (7) and finally, the characteristic moment capacity mc of the beam strengthened with a tensile uhpfrc layer is mc = ζ(fykas + ftkbtγ) (8) 192 vol. 36/2022 reliability of rc beams strengthened with uhpfrc as b d c′ tγ uhpfrc layer h c na εst εfs εcc 0.85fc fstas ft a fcc fst ffs a 0.85fc fcc ffs + fst c ζ z 0.85fc figure 2. stress block geometries for the for the proposed approach considered with tensile capacity of uhpfrc layer due to steel fibres considered 3.2. comparison of proposed approach to test results 12 rc test beams strengthened with tensile uhpfrc layers from previous experimental protocols [3, 12, 13] are used to validate the proposed stress-block approach. figure 3 shows a comparison in moment capacity prediction between the proposed analytical procedure and what was experimentally reported. a significant correlation between the proposed approach and test results is observed. 10 20 30 40 50 60 70 80 0 10 20 30 40 50 60 70 80 figure 3. comparison of a proposed analytical procedure with experimentally predicted moment capacity. 4. reliability assessment of rc beams strengthened with uhpfrc 4.1. basic concepts the current set of recommendations for the assessment of existing concrete structures given by the iso 13822:2010 [14], probabilistic model code by the joint committee on structural safety (jcss pmc 2010 [15]), and the iso 2394:2010 [16] are strongly based on reliability concepts. the fib bulletin 80 [17] gives a codified reliability-based framework that is consistent with iso 2394:1998 [18] for the upgrading of existing concrete structures. generally, reliability targets of βup,50 = 3.3 and β0,50 = 2.3 are given in fib bulletin 80 [17] for upgrading and assessing existing structures related to the ultimate limit states for reliability class 2 and 50 years reference period, respectively. for a reference period of 1 year, these values correspond to reliability indices of about βup,1 = 4.3 and β0,1 = 3.5, respectively. considering the ultimate limit state of flexure, for a rc beam strengthened with a uhpfrc tensile layer and originally designed to take an annual peak moment mapp, the performance function g required for reliability analysis can therefore be cast as g = θmc − ke mapp (9) where θ and ke are the uncertainties in moment capacity and applied moment respectively (see table 1). statistical properties of θ were assessed by comparing the present stress block approach with the test results in figure 3. using similar reliability formulations as stated in [27], annual failure probability pf,1 and reliability index β1 can readily be computed by first order reliability method (form) and monte carlo simulation (mcs). 193 l. simwanda, n. de koker, c. viljoen, aj. babafemi acta polytechnica ctu proceedings parameter statistical model char mean cov ref. b, mm gaussian 150 152.50 0.01 [19] h , mm gaussian 250 258.20 0.02 [20] 1)as, gaussian as 1.021as 0.0125 [21] ef s, gpa lognormal 200 202 0.01 [21] fy , mpa lognormal 500 573.15 0.083 [20] fc, mpa lognormal 39.50 52.25 0.17 [22] ft, mpa lognormal 12 19.66 0.3 [22] ρc, kg/m3 gaussian 2268 2315 0.012 [21] 2)qk, kn gumbel 0.682qk 0.25 [23] c′, mm gamma 25 26.20′ 0.03 [23] l, m constant 2.2 [23] 3)tγ, mm constant tγ [23] τf , mm lognormal 4.15 4.72 0.06 [24, 25] lf /df , m constant 6.5 θ, [-] lognormal 1.08 0.16 ke , [-] lognormal 1.01 0.1 [26] notes: 1) the values of as are updated using the stochastic model in figure 1. 2) qk is adjusted for each beam so that the strengthened beam is critical when analysed. and designed via the present stress block approach. 3) tγ is varied to determine its variation with β1 (section 4.3). table 1. statistics of stochastic parameters in the load, geometric, and resistance models 4.2. reference beam and data collection the rc beam strengthened with a tensile uhpfrc layer to be considered in this work is based on a previous study [4]. the cross sectional dimensions of the beam were 150 mm by 250 mm, with a span of 2.2m. the rc beam was strengthened with 50mm thick layer of uhpfrc on the tensile side. the beam was reinforced with two 500mpa high strength steel rebars with a diameter of 12mm, at a cover of 25mm. the 28 days characteristic compressive strength of the beam and of the uhpfrc layer were 39.5 mpa and 164 mpa, respectively. the average tensile strength of uhpfrc was reported as 12 mpa. to account for the uncertainty in each of the parameters of the beam describing the performance function, a stochastic description of each parameters in terms of probability distribution, mean value and coefficients of variation were defined based on previous literature (table 1). the rc beam strengthened with the uhpfrc tensile layer is analysed and designed with characteristic imposed load qk adjusted for the beam to be critical. specifically, the characteristic imposed load needed to obtain the strengthened beam via the present stress block approach is determined, after which the corresponding mean imposed load is obtained using the assumed coefficient of variation. in addition, the dead load is considered as the self-weight of the beam with its uncertainty accounted for via uncertainties in the geometric properties of the strengthen beam and density of concrete. 5 10 15 20 25 30 35 0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 figure 4. required upgrading reliability increase β1,upgrade as a function of the thickness of the uhpfrc layer tγ and the year when the strengthening operation is performed. 4.3. data analysis to apply the proposed methodology for assessing and upgrading the reliability of rc beams strengthened with uhpfrc layers, lets consider a rc beam under the effect of corrosion whose annual reliability index at year n has been assessed and say deteriorated to βup,1 = 4.3. techniques presented in [28] may be applied for this purpose. the evolution of the reliability 194 vol. 36/2022 reliability of rc beams strengthened with uhpfrc 10 20 30 40 50 3 3.2 3.4 3.6 3.8 4 4.2 4.4 4.6 4.8 5 figure 5. evolution of the reliability index of the rc beam prior to strengthening strengthening (blue solid line) and after addition of 14mm thick uhpfrc tensile layer at 28 years old (black dotten line). index of the rc beam prior to strengthening up to the time β1 reaches a value of βup,1 = 4.3 is given by the blue curve in figure 4. now, regarding the requirement of upgrading this rc beam back to βtarget,1 = 4.7 by strengthening with uhpfrc tensile layers, the thickness of the uhpfrc layer tγ that guarantees a reliability upgrade (β1,upgrade = βtarget,1 − βup,1 = 4.7−4.3 = 0.4) needs to be determined. by computing the increase in reliability index for increasing tγ of the rc beam considered to be upgraded at year n, figure 4 was obtained and may be used for estimating tγ given the required increment in the reliability (∆β). assuming continuation of the corrosion process, the reduction of the reliability index after the strengthening operation can be computed to estimate the remaining service life of the beam (shown by a dotted black curve in figure 5). to decipher the relative importance of the stochastic basic variables on the computed reliability indices, a form-based sensitivity analysis is conducted (see figure 6). the live load q has the highest influence on the computed β values followed by the model uncertainties in the resistance and load models, respectively. as expected, the variables with the lowest variability (table 1) have insignificant influence of the computed β values. most importantly, the α value of model error θ serves as an indication of the required partial resistance factors for the stress-block formulation presented in this work. detailed investigation on the statistical characteristics of θ based on a larger database remains an avenue for future studies q a s f y h c b f c f t k e 0 0.2 0.4 0.6 0.8 1 figure 6. first order reliability method (form) sensitivity factors of the stochastic basic variables. deterministic basic variables have form α values of zero. 5. conclusions a methodology for assessing and upgrading the reliability of rc beams strengthened with uhpfrc tensile layers has been presented. the applied methodology includes stochastic stress-block analysis of a section, assuming a perfect bond between the rc beam and the uhpfrc layer, and then conducting life-time reliability of the rc beam before and after the strengthening operation. deterioration induced by corrosion can be incorporated into the analysis via stochastic corrosion models. a plot for determining the required thickness of the uhpfrc tensile layer to upgrade to the required reliability level is also given, considering the time from construction when the strengthening operation is conducted. 195 l. simwanda, n. de koker, c. viljoen, aj. babafemi acta polytechnica ctu proceedings the present work has only considered the timedependent deterioration model in terms of the corrosion of steel reinforcing area. consideration of timedependent behaviour of concrete, uhpfrc and steel strength, other geometric properties of concrete and the uhpfrc layer, and the time-dependent correlations in reliability analyses should be considered in future studies, with more detailed models and availability of information on the time-dependent correlation of deteriorating rc beams strengthened with uhpfrc layers. 6. list of symbols a, depth of a rectangular stress block as, area of steel rebar in tension b, width of the beam c, depth to the neutral axis d, depth from extreme compressive fiber to centroid of rebar steel ef s, modulus of elasticity of fibers fc, compression strength of uhpc ft, the tensile stress of uhpc fy , yield strength of steel rebar fst, stress in steel rebar ρc, density of concrete h, height of the beam c′ cover to steel rebar lf , length of fibers df , diameter of the fibers ρs, rebar percentage σf y , fiber yielding stress τf , fiber-concrete bond strength β1, stress block parameter ρc, density of uhpfrc ρst, density of steel vf , percentage fibre content tγ, thickness of the uhpfrc layer. acknowledgements reliability analysis computations were performed using the uqlab reliability analysis library [29]. references [1] t. noshiravani, e. brühwiler. analytical model for predicting response and flexure-shear resistance of composite beams combining reinforced ultrahigh performance fiber-reinforced concrete and reinforced concrete. journal of structural engineering 140(6):04014012, 2014. https://doi.org/10.1061/(asce)st.1943-541x.0000902. 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[29] s. marelli, b. sudret. uqlab: a framework for uncertainty quantification in matlab. in vulnerability, uncertainty, and risk, pp. 2554–2563. american society of civil engineers, reston, va, 2014. https://doi.org/10.1061/9780784413609.257. 197 https://doi.org/10.1007/s11367-011-0297-3 https://doi.org/10.35789/fib.bull.0080 https://doi.org/10.1007/s11367-011-0297-3 https://doi.org/10.1016/j.engstruct.2013.07.015 https://doi.org/10.1111/j.1559-3584.2002.tb00128.x https://doi.org/10.1016/j.cemconcomp.2017.02.006 https://doi.org/10.1016/s0140-6736(06)69720-1 https://doi.org/10.6028/nbs.sp.400-64 https://doi.org/10.1016/j.engstruct.2006.02.005 https://doi.org/10.1080/15732479.2016.1198394 https://doi.org/10.1016/j.engstruct.2021.112767 https://doi.org/10.1016/j.strusafe.2007.02.002 https://doi.org/10.1061/9780784413609.257 acta polytechnica ctu proceedings 36:191–197, 2022 1 introduction 2 stochastic modelling of deterioration of rc beams prior to strengthening 2.1 corrosion modelling 2.2 corrosion stochastic simulation 3 analysis of rc beams strengthened with uhpfrc 3.1 moment capacity: proposed approach 3.2 comparison of proposed approach to test results 4 reliability assessment of rc beams strengthened with uhpfrc 4.1 basic concepts 4.2 reference beam and data collection 4.3 data analysis 5 conclusions 6 list of symbols acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.37.0031 acta polytechnica ctu proceedings 37:31–37, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague proposal of point-wise power reconstruction for irt-4m fuel assemblies in coupled monte-carlo calculations pavel jíška∗, jan frýbort czech technical university in prague, faculty of nuclear sciences and physical engineering, department of nuclear reactors, v holešovičkách 2, 180 00 prague 8, czech republic ∗ corresponding author: jiskapav@fjfi.cvut.cz abstract. current efforts to develop high-fidelity multi-physics tools for research reactors with irt-4m fuel leads to the involvement of monte-carlo transport and subchannel thermo-hydraulic codes in the process of safety assessment. monte-carlo based neutronic codes are often used as a reference tool and are suitable for a prediction of a detailed power distribution. this paper proposes a method of the point-wise power reconstruction based on a radial basis function (rbf) interpolation for complicated irregular geometries and compares different detector functionalities of the serpent2 code from the time efficiency point of view. furthermore, a preliminary validation of the applied method for irt-4m in a subchannel code subsals mesh is presented. this code is currently under the development at the research centre řež and the czech technical university in prague and focuses on the thermohydraulic analysis of irt type fuel assemblies. keywords: power reconstruction, multi-physics coupling, irt-4m, serpent2, subsals. 1. introduction the main motivation behind the development and the application of advanced calculation tools and methods is the ensurement of ability to simulate the processes in the reactor core with a lower degree of uncertainity than nowadays computing resources allow, and even beyond operational safety analyses. most of the attention is predominantly given to power reactors. however, research reactors can also have difficulties to simulate specific phenomena. two czech research reactors lvr-15 [1] and vr-1 [2] are operated with irt-4m fuel assemblies. this type of fuel has an unusual irregular geometry as demostrated by figures 1a or 1b. its basic description is provided by table 1. because of its unique concentric square annular fuel elements with rounded corners, the coolant flow is subjected to a relatively significant pressure driven crossflow. a detailed calculation of such thermalhydraulic characteristics is beyond capabilities of standard dedicated one-dimensional system codes. a new subchannel code subsals (subchannel square annular layout solver) is under the development at the research centre řež and the czech technical university in prague. the code is currently in the phase of concept verification. nevertheless, its further utilisations are considered, for instance, its coupling with neutronic codes. monte-carlo based neutronic codes are often used as a reference tool and are suitable for prediction of a detailed power distribution. for multi-physical coupling, it is essential to ensure the appropriate data transfer between individual codes in a required data designation value fuel assembly type irt-4m standard number of fuel tubes in fa 8 fuel meat material uo2 + al enrichment 19.7 % cladding material sav-1 alloy mass of 235u in fa 300 g core pitch 71.5 mm cladding thickness 0.45 mm fuel meat thickness 0.70 mm water gap between fuel tubes 1.85 mm active height of fuel meat 600 mm outer dimensions of tubes and corner radius: 1st tube 2nd tube 3rd tube 4th tube 5th tube 6th tube 7th tube 8th tube (cylindrical) 69.6 mm; 9.3 mm 62.7 mm; 8.5 mm 55.8 mm; 7.7 mm 48.9 mm; 6.9 mm 42.0 mm; 6.1 mm 35.1 mm; 5.3 mm 28.2 mm; 4.5 mm 21.3 mm table 1. basic characteristics of standard irt-4m fuel assembly [3]. structure. different embedded funcionalities for power reconstruction (i.e. different detector types) are implemented in monte-carlo code serpent2 [4]. however, some approaches appear to be more computationally challenging or difficult to implement in such a complex irt-4m geometry. fuel assembly power mapping can be regarded as an interpolation problem of scattered data which is common in many fields. the radial basis function interpolation offers excellent tool for interpolating multidimensional scattered data whose value is dependant only on system coordinates. thus, it is an 31 https://doi.org/10.14311/app.2022.37.0031 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en pavel jíška, jan frýbort acta polytechnica ctu proceedings (a) . mesh 9 × 9 (b) . mesh 17 × 17 (c) . subsals mesh figure 1. cartesian mesh detectors 9 × 9 and 17 × 17, and subsals thermal-hydraulic model for 8-tube irt-4m fuel assembly. appropriate approach for a power and temperature mapping in situations where values of scattered variables varies gently. in addition, the round-off error of a piecewise constant distribution could be minimized by a continuous function. however, it should be noted that the application of rbf interpolation is prone to uncertainty of sampled data, which may apply especially for monte-carlo tallies. 2. theory within this work two different types of response detector meshes were considered for power reconstruction in the irt-4m fuel assembly. those meshes are denoted here as: (1.) mesh of cell detectors (mocd) is a general mesh made of individual cell detectors (see option dc in [5]) situated only in the fuelled regions in accordance with required subsals layout (see figure 1c), (2.) irregular cartesian mesh detector (icmd) is a standard serpent2 irregular cartesian mesh (see option dmesh with datamesh type 6 in [5]). mocd enables to determine values of a variable in a general spatial mesh superimposed on the base area of an fuel assembly, whereas irmd can be applicable only to an irregular cartesian grid problem. therefore mocd is here considered as a reference mesh for point-wise rbf-based power reconstruction testing. the icmd is here utilized as an auxiliary grid on which the mean values of an intensive variable (i.e. power density) are interpolated by rbf interpolation. here should be noted, that the term point-wise power reconstruction is here used to distinquish from a frequently used term pin-wise or plate-wise power reconstruction since the process of evaluation is different and the final value is not connected with any fuel pins or plates, but rather with a general volume and its center of mass. basic descriptions of mesh and cell detectors are mentioned in this section. further, basic information about rbf interpolation for the pin-wise power reconstruction in irt-4m fuel assemblies are provided. 2.1. mesh/cell detectors monte-carlo mesh detectors are able to determine various integral quantities in user-defined detector bins. the fundamental idea is based on a division of space into partial regions and a selection of events of particular interest. the resultant spatial description of integral quantities has the character of a piecewise constant function. for fission energy production p is evaluated integral of type p = ∫ v ∫ e κσf (r, e) ϕ(r, e) d3r de, (1) where κσf (r, e) is macroscopic total fission energy production cross section and ϕ(r, e) is neutron flux density [6]. in order to achieve the highest possible level of statistical accuracy, it is reasonable to increase the size of cells in mesh detectors or to increase the number of neutron tallies. on the contrary, in an effort to characterize the distribution function as accurately as possible, it is necessary to minimize the rounding error of the mesh detector approximation. however, the common limit condition is the computational performance. for any simulation, it is necessary to find the balance between the level of the detail and the simulation time [7]. 2.2. radial basis function theory radial basis functions interpolation is one of the commonly used methods to interpolate multi-dimensional data and have gained popularity over the last few decades in a variety of areas such as a surface reconstruction, multivariate interpolation, approximation theory, neural networks, machine learning, etc. rbfs are typically used to approximate an unknown func32 vol. 37/2022 tion f : rn → r by the continuous function s(x) = n∑ i=1 λi φ(||x − xi||), (2) where {xi, i = 1, 2, ..., n } ∈ rn are given scattered data coordinates, φ : ⟨0, ∞) → r denotes a radial basis function, and ||x − xi|| is the euclidean norm between coordinates x and xi [8]. the approximating function, to be constructed over n sampled coordinates, satisfies s(xi) = fi, i = 1, 2, ..., n, (3) where values fi of unknown function f are assumed to fulfill fi = f (xi). the equation (2) could be rewritten in terms of linear system of equations aλ = f , (4) or in matrix form  φ(r1,1) . . . φ(rn,1)... . . . ... φ(r1,n ) . . . φ(rn,n )     λ1... λn   =   f1... fn   , (5) where rj,i = ||xj − xi||. from the definition of the euclidean norm, it is clear that matrix a is symmetric with dimensions n × n . to ensure a unique solution, the matrix a has to be non-singular. specific form of the radial basis function φ(r) may have a different prescription (e.g. gaussian, multiquadric, polyharmonic, etc.). 3. power reconstruction currently, the serpent2 program focuses predominantly on reactor systems with regular grids, similar to standard power reactors. its coupling interface is not fully adapted to generally irregular geometries, as required in the subsals code for irt-4m representation. however, this minor deficiency can be easily circumvented using a specific type of mesh of cell detectors and an individual coupling wrapper. in the proposed subchannel model, for the irt-4m fuel assembly in subsals program (see figure 1c), three different types of heat sources occur: • one-eighth annular section, • large straight plate, • small straight plate. 3.1. models description serpent2 models includes simplified three-dimensional lvr-15 reactor core configuration with standard (8tube) fresh fuel assemblies irt-4m surrounded by beryllium blocks (see figure 2). depending on the level of detail of fuel meat division inside one preselected fuel assembly, two geometrically different models can be distinguished: (1.) simplified fa model where all fuel tube cells and water channels are considered as an unique material cell circumscribed by square prisms with rounded corners, (2.) detailed fa model which is almost identical to the simplified fa model with the only exception in the one core position (b6 or e6), where fuel meat materials are replaced by the identical material azimuthally divided into the subsals computational mesh (i.e. 8 fuel meat cells replaced by 168 cells). figure 2. serpent2 model with irt-4m fuel assemblies surrounded by beryllium blocks (framed positions b6 and e6 of particular interest). depending on the implementation of response detector meshes (see mocd and icmd), three different approaches were considered for determining the point-wise power distribution in the subsals layout. the nomenclature for each approach is stated below. these notations will be used throughout this work for a simplification: (1.) cell-based model where the detailed fa model is utilized and a mesh of cell detectors (mocd) is directly included in the main universe as individual fuel meat cells, (2.) universe-based model where the simplified fa model is utilized and a mesh of cell detectors (mocd) is included in a separate universe as mesh of individual detector cells overlapping the main universe, (3.) rbf-based model where the simplified fa model is utilized and an irregular cartesian mesh detector (icmd) is applied 33 pavel jíška, jan frýbort acta polytechnica ctu proceedings (a) . position e6 (b) . position b6 figure 3. reference power density imbalance factors k(i)q in the subsals mesh. for a subsequent two-dimensional rbf interpolation.1 3.2. testing settings all results were generated by running serpent2 (version 2.1.32) in a criticality source simulation mode. the run parameters were set with 40 inactive cycles to populate the fission source, 400 active cycles to collect results, and 500 000 neutrons per cycle. all calculations were performed on one computer using 32 openmp threads with processor amd epyc 7351. the host operating system was an ubuntu 20.04.4 lts. 3.3. implementation of rbf-based power reconstruction it is necessary to mention at the outset that tested rbf-based power reconstruction method utilizing rbf interpolation is established on rough simplifying assumptions. the method is based on the known distribution function in the auxiliary grid (e.g. icmd). the auxilliary mesh selection for irt-4m fuel assembly is considerably complicated task and can have a significant effect on reconstruction outputs. source data for the rbf-based power reconstruction are total values of power in auxiliary cartesian mesh (i.e. icmd), or respectively mean power density if related to the relevant fuel meat volume. however, the question how to assign function values (i.e. power densities) to the spatial coordinate remains. usually they are assigned to the center of mass of given fuel meat element (e.g. in pin-wise reconstruction). nevertheless, the center of mass of icmd cell may not coincide with the center of mass of the fuel meat cell. unfortunately, this may result in the incorrect assignation of function values (especially when multiple fuel cells are in cartesian mesh cell or when rounded corners are included) and thus resultant interpolant could be disorted. in this work, the mean values were 1please note, that the manner of response detector mesh implementation is external for universe-based and rbf-based models, whereas internal for a cell-based model. always assigned to the center of mass of cartesian mesh cell detector. the common aim was to create cartesian mesh so that center of masses matches (at least in planar segments). another assumption is that resulting power distribution is obtained by quantifying the interpolation function at predefined coordinates and by volumetric multiplication of given fuel meat element. predefined coordinates were chosen manually to correspond approximately to the center of mass of fuel meat element in subsals layout. however, the assumption is fulfilled only if the course of the function is not convex or concave in the given interval. within this work, the polyharmonic radial basis function φ(r) = r3 (6) was selected for testing. the choice of rbf was based on the good practice and meaningful reconstruction outputs. the utilisation of other radial basis function is possible, but their impact on results was not evaluated here. an auxiliary program for the rbf-based power reconstruction has been created. the program utilizes python3 language and is based on numpy and scipy libraries (version 1.23.1, respectively 1.9.0). the program has an automatic procedure, which can evaluate irt-4m fuel meat volumes almost in any icmd cell to ensure proper volume division and also includes plotter in subsals mesh (see figure 1c) for result visualisation such as figures 3. 3.4. testing rbf-based power reconstruction the point-wise rbf based power reconstruction was tested on two different auxiliary icmd, specifically 9 × 9 and 17 × 17. meshes were set as demonstrated on figures 1a and 1b. the division of the detector in the vertical direction was not considered. performed point-wise rbf based power reconstructions were compared against the reference data obtained directly in subsals layout (i.e. direct comparison with universe-based model results). it was 34 vol. 37/2022 (a) . position e6 (b) . position b6 figure 4. rbf interpolation of reference data. ensured, by the calculation settings, that the statistical uncertainity for individual fuel meat elements was below 0.5 % and the case of cartesian meshes for individual cells up to 1 % (especially for cells with lowest contain of fuel meat, i.e. cells (4,4) and (5,5) in mesh 17 × 17). in the figure 3, reference values of power density imbalance factors k(i)q for fuel tubes cells in core positions e6 and b6 are depicted. power density imbalance factors can be expressed by equation: k(i)q = q (i) v qv , (7) where q(i)v is a power density in cell i and qv denotes a mean fuel assembly power density in given axial layer. in the figure 4 are demonstrated rbf interpolations of reference data in positions e6 and b6. when comparing point-wise rbf-based power reconstructions for both cartesian meshes with reference data (see figures 5 and 6), it is possible to quantify relative deviations. those are summarized by the table 2. overall, presented and roughly simplified rbf based power reconstruction gives reasonable results, because the maximal fission power density difference between neigbouring fuel meat elements could reach roughly 13 % for fuel assembly in the b6 possition and 23 % in the e6. the round-off error of irregular mesh reconstruction is higher than the error of point-wise rbf based power reconstruction. position and icmd max.deviation rmse b6 mesh 9 × 9 3.29 % 1.15 % e6 mesh 9 × 9 4.29 % 1.28 % b6 mesh 17 × 17 2.33 % 0.91 % e6 mesh 17 × 17 3.51 % 1.03 % table 2. max. relative differences of rbf-based power reconstruction from reference data. figure 5. relative differences between rbf-based power reconstruction and reference data for cartesian meshes 9 × 9 (top) and 17 × 17 (bottom), both for fuel assembly in e6. 4. computational speed the computational performance of the monte-carlo simulation depends predominantly on the complexity of created model. the total number of cells and their boundary surfaces play an important role in this case. creating strongly heterogeneous discrete models (i.e. models whose properties are characterized by many 35 pavel jíška, jan frýbort acta polytechnica ctu proceedings parameter simplified fa w/o detectors detailed fa w/o detectors rbf-based model cell-based model universe-based model9 × 9 17 × 17 total cpu time [min] 2358.83 2694.85 2378.52 2522.38 5745.35 7721.40 total wall-clock running time [min] 92.74 105.12 92.73 95.19 197.83 260.95 wall-clock time for transport simulation [min] 92.00 104.50 92.00 94.52 197.18 260.33 cpu usage fraction in transport simulation [–] 25.70 26.10 25.90 26.85 29.23 29.78 allocated memory size [mb] 7586.46 7589.86 7586.46 7586.46 7590.06 7588.57 memory used for storing cross sections [mb] 2809.81 2809.81 2809.81 2809.81 2809.81 2809.81 memory used for storing material-wise data [mb] 1086.42 1086.42 1086.42 1086.42 1086.42 1086.42 total number of cells [–] 3305 3465 3305 3305 3465 3473 total number of cells using unions [–] 0 0 0 0 0 0 table 3. comparison of serpent2 simulation characteristics for different cases. figure 6. relative differences between rbf-based power reconstruction and reference data for cartesian meshes 9 × 9 (top) and 17 × 17 (bottom), both for fuel assembly in b6. materials and a common, usually mean, value of state variable) can be not only computationally inefficient, but also user-unfavorable. the comparison of main computational characteristics is provided in table 3. as can be seen, the utilisation of a finer model division of fuel meat in the position b6 leads to an extension of computing time by approximately 14 % (from the dirrect comparison of simplified fa and detailed fa model without any detectors). if other fuel meat elements were divided, for instance in other fuel assemblies, the effect of the time increase will be even more significant. consequently, performing full core calculations with cell detectors can be considerably time disadvantageous. in terms of the pin-wise power distribution reconstruction, a significant increase of computational time was observed in the case of cell-based and universebased models. from table 3 is evident that the utilization of cell detectors has considerable impact on the transport simulation time. in direct comparison with rbf-based models (i.e. standard icmd), this increase can be quantified for cell-based model by 144 % and for universe-based model even by 224 %. in both cases, an increase of fraction of time spent in openmp parallel loops was observed (by 15 % and 19 %, respectively). 5. discussion modification of the cell detectors implementation in monte-carlo code serpent2 may allow to detect local neutron-physical characteristics more efficiently, with lower computing time and a more user-friendly model definition (i.e. without any update of model, which also applies to models considering detailed burnup distributions). to focus on the cell of particular interest, which is not included in geometry, could be crucial in neutron-physical and thermal-hydraulic coupling. in that case, several iterations (or criticality source simulations) are performed until the set convergence criterion is met. observed time increase will be multiplied, and even more if simulating full core coupled calculations. however, the author is not fully acquainted with the implementation of cell detectors in the serpent2 code. this could be further discussed with code developers. if the modification is not feasible, a simplificated method of rbf-based power reconstruction for irt4m fuel assembly was presented to shorten the computational time. for application of this method, the reduction of statistical uncertainty is necessary and ensurement of proper values to coordinates pairing. further optimization of this method is possible as demostrated in the section 3.4. furthermore, this method could be applicable to the temperature characteristics of fuel and coolant materials in coupled 36 vol. 37/2022 calculations. the pleasant benefit is that coupled monte-carlo models could not have to be geometrically complicated (not necessary to use a discrete material segmentation) and potentially save computinal time spent in a neutron transport simulation. the local temperatures of each fuel meat element, or coolant channel, could be also described by rbf interpolation (i.e. one function prescription) for given material region. this may result in a potential reduction of computing time in coupled calculations, because of less number of cells in transport calculation. however, this temperature field interpolation has not been computationally verified so far. 6. conclusions in this work a new method for power reconstruction for irt-4m fuel assemblies was presented. the rbf-based power reconstruction method was tested to provide quantitative characterization of method error and the computational demands. testing was performed on simplified lvr-15 reactor core for two selected fuel assembly positions b6 and e6 with different assembly-wise power distribution. demonstrated application gives reasonable results with errors lower than a round-off error of neighbouring fuel element cells. although rough simplifications were used, the maximal observed deviation was approximately 4.3 %. maximal power density imbalance rmse for fuel assembly reaches less then 1.3 %. maximal deviations decrease with finer division of mesh. nevertheless, with finer mesh cells, larger local uncertainties may appear. attention need to be paid, when choosing input data from a cartesian data mesh. as a part of this work was found, that cell detectors implemented in monte-carlo code serpent2 (version 2.1.32) has a significant impact on computational time. when using cell-based model for direct power reconstruction (see section 3.1), an increase in computing time around 144 % was observed. if universe-based model (see section 3.1) was used, the time extension was approx. 227 %. this outcome is going to be further discussed with code developers. the new update of cell detectors of serpent2 may allow to detect local neutron-physical characteristics more efficiently, with lower computing time and a more userfriendly model definition (i.e. without any update of model). list of symbols a matrix of evaluated kernels cp u central processing unit e energy [j] fi function value in xi icm d irregular cartesian mesh detector κ recoverable energy per fission [j] k (i) q power density imbalance factor for node i [–] λi linear combination coefficient max maximum m ocd mesh of cell detectors p total power [w] q (i) v power density in node i [w m−3] qv mean fuel assembly power density [w m−3] r position vector ||r|| euclidean norm rbf radial basis function rm se root mean square error σf fission macroscopic cross section [m−1] s(x) aproximating radial basis function interpolation ϕ neutron flux [m−2 s−1] φ radial basis function kernel v volume [m3] x coordinates x-axis [cm] y coordinates y-axis [cm] acknowledgements this work was supported by the grant agency of the czech technical university in prague, grant no. sgs22/189/ohk4/3t/14. references [1] research centre řež. research reactor lvr-15. http://reaktory.cvrez.cz/en/ research-reactor-lvr-15/, 2022. [2022-10-10]. [2] l. frýbortová, j. rataj, l. sklenka, et al. the training reactor vr-1 – 30 years of operation. 21st international scientific conference on electric power engineering (epe) pp. 1–7, 2020. https://doi.org/10.1109/epe51172.2020.9269254. [3] e. vilímová, p. dostál. provozní bezpečnostní zpráva reaktoru lvr-15. research centre řež, 3rd edn., 2021. [4] j. leppänen, m. pusa, t. viitanen, et al. the serpent monte carlo code: status, development and applications in 2013. annals of nuclear energy 82:142–150, 2015. https://doi.org/10.1016/j.anucene.2014.08.024. [5] vtt technical research centre of finland. input syntax manual – serpent wiki. https://serpent.vtt. fi/mediawiki/index.php/input_syntax_manual, 2022. [2022-10-10]. [6] j. leppänen. development of a new monte carlo reactor physics code. vtt technical research centre of finland, 2007. [7] d. p. griesheimer. functional expansion tallies for monte carlo simulations. ph.d. thesis, university of michigan, 2005. [8] g. song, j. riddle, g. e. fasshauer, f. j. hickernell. multivariate interpolation with increasingly flat radial basis functions of finite smoothness. advances in computational mathematics 36(3):485–501, 2012. https://doi.org/10.1007/s10444-011-9192-5. 37 http://reaktory.cvrez.cz/en/research-reactor-lvr-15/ http://reaktory.cvrez.cz/en/research-reactor-lvr-15/ https://doi.org/10.1109/epe51172.2020.9269254 https://doi.org/10.1016/j.anucene.2014.08.024 https://serpent.vtt.fi/mediawiki/index.php/input_syntax_manual https://serpent.vtt.fi/mediawiki/index.php/input_syntax_manual https://doi.org/10.1007/s10444-011-9192-5 acta polytechnica ctu proceedings 37:31–37, 2022 1 introduction 2 theory 2.1 mesh/cell detectors 2.2 radial basis function theory 3 power reconstruction 3.1 models description 3.2 testing settings 3.3 implementation of rbf-based power reconstruction 3.4 testing rbf-based power reconstruction 4 computational speed 5 discussion 6 conclusions list of symbols acknowledgements references my title https://doi.org/10.14311/app.2022.36.0224 acta polytechnica ctu proceedings 36:224–230, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague the role of multi-fidelity modelling in adaptation and recovery of engineering systems rui teixeiraa, ∗, beatriz martinez-pastora, maria nogalb, alexandra micuc, alan o’connorc a university college dublin, school of civil engineering, a richview newstead belfield, dublin 4, ireland b technical university delft, department of materials, mechanics, management & design, stevinweg 1, 2628 cn delft, the netherlands c trinity college dublin, school of civil, structural and environmental engineering, college green, dublin 2, ireland ∗ corresponding author: rui.teixeira@ucd.ie abstract. significant research has been conducted in identifying optimal recovery and adaptation decisions in disruptive scenarios using engineering models. in this context, an aspect that has been target of limited research is that of response times. modelling is expected to grow progressively more complex as it becomes more accurate. such complexity increases modelling efforts, and the promise of optimal adaptation and recovery may become hindered. the present work discusses the role of modelling fidelities in adaptation and recovery of systems, and in particular that of using a lower fidelity model that enables zero-time analyses of a system. a framework is proposed for using different fidelities in adaptation and recovery, considering system’s decision time requirements. the relevance of this analysis is researched in two traffic networks and results show that multi-fidelity models should be expected to play a key role in increasing the efficiency of optimal adaptation and recovery decisions. keywords: metamodels, multi-fidelity model, system adaptation, system recovery, systems. 1. introduction adaptation and response of civil engineering systems has become one of the most prominent topics of research in recent years. several methodologies and approaches have been proposed to tackle the need for adaptation and responsiveness, however, most of them fail to discuss one of the most relevant aspects of the system’s responsiveness, the decision time. engineering systems are complex in nature. they depend on multiple variables that can interact together and that originates a need for these to be treated as holistic. it is with no surprise that system analysis is recurrently highly complex, relying on intensive data analytics or costly modelling techniques. due to this optimal decision-making for civil engineering systems can become a resource intensive task. most approaches for adaptation and response to natural or man-made disasters lay their foundation on the premise that under a disruption an optimal adaptation decision making scheme can be found that will maintain the functionality of a system at certain minimum operational level. this can be achieved through intervention, by adapting the system operational variables, or other alternative system approaches [1]. in all cases, some form of technique needs to inform decision making, and current trends of research indicate that there is a growing interest in the application of high-fidelity (hf) modelling as a tool to inform decisions. this is identified in the progressive interest for hf tools such as, finite-element methods, and approaches, such as digital twinning or prediction of complex adaptivity in systems. most of these techniques were introduced in order to enable prediction and decision-making that is as close as possible to real-time response. however, in practice, running a decision-making scheme in such hf models can be effort consuming. in the instance of needing to use real-time information and prediction of the system operation to adapt and prevent any perturbation a form of parametric analysis needs to be performed. as this procedure will still depend on the hf model built to have a accuracy on predicting the system’s response, then performing any analysis that will establish a new point of operation that maintains (or optimizes) its functionality is expected to have large cost. within this context, the present paper discusses the role that multi-fidelity (mf) techniques may have in efficient adaptation of engineering systems. mf, that is, a blend of hf and low-fidelity (lf) modelling, may allow to balance the effort that is required to set optimal adaptation and recovery decisions. in particular, decision-times are expected to have a significant influence on the efficiency of the decision-making schemes. in the case of adaptation and recovery of an engineering system to a disruptive event, stalling a decision for a time longer than strictly necessary may set the boundary between significant or minor impacts in the system. in order to discuss the role of mf models in the context described, section 2 introduces the idea of mf modelling, section 3 discusses a framework 224 https://doi.org/10.14311/app.2022.36.0224 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 36/2022 role of multi-fidelity in system adaptation and recovery figure 1. multi-fidelity, from low-fidelity to high-fidelity models [2]. of application to adaptation and recovery of systems, section 4 presents two examples of application to a traffic network, and finally, section 5 draws the main conclusions of the work developed. 2. multi-fidelity (mf) modelling mf modelling has been widely researched in order to tackle the increasing analysis costs that are imposed by progressively more complex and accurate modelling techniques. the fundamental idea of mf is to combine hf models (commonly characterized as of high-resolution, high analysis cost), with lf models (commonly characterized as of lower resolution, low computational cost), see figure 1. hf models are expected to have better accuracy in practice, however, this does not mean that lf models cannot achieve adequate accuracy. in mf, both hf and lf are applied to balance the amount of effort required to perform an analysis based on the principle that an adequate amount of hf data may suffice an accurate analysis if complemented by lf data. such usage of different levels of fidelity has been previously researched and identified to have an important role in fields that are characterised by large analysis efforts such as, uncertainty quantification and optimization [3, 4]. it should be expected to become relevant in all engineering analyses that depend on effort consuming models, and where the time necessary to inform decision-making schemes has a key role. 2.1. forms of multi-fidelity modelling mf models can appear in different forms, and the level of fidelity depends on a series of factors. it can be said that hf models represent the behaviour of a system with an adequate level of accuracy, and that lf models are frequently a less refined representation of these. lf models can be created by assuming simplified forms for hf models. this can be achieved through dimensionality reduction, linearisation, simpler physics models, coarser domains, partially converged results, among other [5]. [5] highlight also different forms in which a system can be distinguished in terms of levels of fidelity. these may depend on the physical assumptions (e.g. may use different theories), numerical assumptions (e.g. more refined grids or different computational models that use same underlying theory), and also on the simple nature of the information (e.g., experimental versus simulations). one alternative that has captivated interest in the field of mf modelling is metamodelling. it consists in creating surrogates of a function or model by characterizing the relationship between inputs and outputs and relies on metamodels as lf pairs to hf data [6]. an overview of popular metamodelling techniques is provided in [7], and [8] provides insight on the usage of mf paired with metamodelling. different metamodels can be applied in this context. the interested reader is directed to the extensive literature on this topic [7] for a review of some examples. in the present work a kriging model is applied to discuss the relevance of mf alternatives in adaptation and recovery of systems. 3. integration of mf modelling in adaptation and response of engineering systems effective system response or adaptation is expected to rely on quick responses to damaging or disruptive events. when an informative state of the system is acquired, if decisions that are based on this state are not implemented fast enough, then this informative state may be lost as the system evolves. moreover, if an optimal decision of recovery or adaptation that was based on this informative state is later obtained; then there is a large likelihood that it won’t be optimal as the system evolved. to add that when an informative state of a system that is being affected by a disruption exists, from the perspective of the decision-maker, it is very difficult to not act and wait for an optimal evaluation of the decision-making scheme, even if this would be the most beneficial choice to achieve efficient decisions. as a result if an optimization for adaptation or recovery is pursued that aims at optimal decision-making, and if the results are delayed by running optimal searches for the decision variables in hf techniques, then loss of efficiency should be expected. in order to illustrate the importance of these aspects, research on mf applications is implemented in a traffic network. two networks are used to research on the effectiveness of mf in contexts of, respectively, recovery and adaptation. the nguyen-dupuis and 225 r. teixeira, b. martinez-pastor, m. nogal et al. acta polytechnica ctu proceedings sioux-falls traffic networks are used. a more detailed description of the network parameters used in these can be found in [9]. the mf model applied combines as hf the traffic network user equilibrium, a nonlinear model that determines the optimum state of equilibrium for the network users; and as lf model a kriging model, a surrogate of the hf model. 4. discussion on recovery and adaptation of a traffic network to a disruptive event the mf relationship between the hf and lf in the contexts of adaptation and recover is highlighted in figure 2. two essential notions of time under decisionmaking schemes can be highlighted in this context: • slow-time: where time is not key for the effectiveness of overall decision-making response. • fast-time: where time has critical role in the overall effectiveness of the decision-making scheme. a metamodeling discussion that uses a similar idea was previously implemented in the research of digital twinning in [10]. figure 2. framework for mf application in adaptation and response. for most of the operation of a system, decisionmaking can be considered to be in slow-time. any modelling of decision-making in such circumstance can rely in hf models and limited change to steady conditions should be expected. in this phase, hf can predict outcomes for the system at limited cost. effectiveness will be largely unaffected by decisiontimes as the pairing of the system modelling with its operation is continuous. however, under a disruptive scenario the same assumptions about the relationship of decision-making schemes and effectiveness are not expected to hold. in unsteady conditions, the system will be changing, hence, an optimal decision on slowtime is unlikely to be optimal (to adapt or recover the system a decision-making scheme may be needed to adapt the decision variables). in such scenario optimal decisions should have zero-time and be paired with the system unsteady conditions. it is within this idea that mf modelling is framed in system recovery and adaptation. mf allows to control time in optimal decision-making, depending on the system demand. in slow-time the multi-fidelity technique can be improved as a predictor of the system response for fast-time. in this phase, which is representative of normal operational conditions, the hf model responds and the lf model improves using hf data (including potential disruptive scenarios). to construct the mf model for this effect, it is assumed that a hf model f (x) exists such that it is an exact approximation of the system behaviour s(x) for any operational state x, then in slow-time a mf model can be built using metamodelling that includes a lf surrogate g(x) of f (x). this lf can use system inputs under foreseen and unforeseen scenarios of disruption or operation, x. if a global approximation to the space of x can be achieved, then the lf metamodel can act as an accurate predictor of f (x) and s(x) for the whole domain of operation. application of the kriging enables lf improvement in slow-time using its properties to enrich the lf model so that it responds accurately in fast-time. in this phase, it is possible to train the lf model in x using hf data. this relies on operational data and exploits the kriging unique gaussian uncertainty. with kriging as a lf, it is possible to search the operational space x for points of large uncertainty. the lf has then the potential to systematically become a more accurate prediction of the hf model. as a result, in fast-time, the lf swaps to the responding position, and the hf is expected to improve collecting data from unforeseen occurrences. in this phase decision-times are critical, and the lf model, i.e. the kriging, is expected to provide a sufficiently appropriate prediction in order to allow for a quasi-optimum (with relation to the lf-hf error) decision-making. moreover, for any x that does not fulfil the accuracy condition in fast-time a hf sample can be taken to enrich the lf model. in fast-time, samples at hf should be minimized and only extracted when necessary. two examples of application, in recovery and adaptation, are studied as reference examples of the relevance of mf in the context described. these are presented in sections 4.1 and 4.2. 4.1. recovery of traffic network in a scenario of damage in the first example of application, that concerns recovery of a traffic network, the nguyen-dupuis network is used. [11] researched this same traffic network in the context of reliability. it is assumed that a damage affects 7 of its links at 50% of their capacity; links 1,7,8,9,10, 12, 23 as per figure 3. then four recovery teams are to be deployed in order to mitigate the effects of this damage and to recover the network 226 vol. 36/2022 role of multi-fidelity in system adaptation and recovery functionality. this recovery will be as efficient as the decision-making that plans the recovery strategy. in practical terms, an efficient recovery should restore the network operational performance to the a stage of pre-damage as soon as possible. such procedure depends on an efficient allocation of resources, which minimizes the loss of functionality in the network and improves its resilience [12, 13]. a damage in the network will increase the cost for users to travel from one node in the network to another, in particular if this trip passes through an affected link. from a decisionmaker perspective the interest is to distribute the resources that will efficiently decrease the travel cost. if a damage occurs in the system at t = 0 then in this moment the decision-maker should set the optimum strategy in place to restore the network so that in t=1 its functionality starts to be recovered. in the present case the lf model is defined using kriging where for x states of operation, the mf model uses the lf surrogate enriched on the hf model (the network user-equilibrium model) in slow-time. x encloses potential link capacities for the considered links. in the seven-dimensional operational state (i.e., of damaged links), only 28 hf evaluations in slowtime are required to set an accurate lf model with a maximum relative error of prediction of a latin hypercube sample (see example [13, 14]) of 100 points of 5%. in order to discuss the relevance of the approach proposed, figure 4 presents the results of recovery using the mf approach. the mf is used in fast-time to predict allocation decisions. the results of this approach are then compared with the result from decision-making that relies exclusively in the hf technique. in both cases a genetic algorithm (ga) is used to allocate teams for recovery. results show that recovery decisions for the mf approximate with small loss of accuracy the results provided by the hf. mf prediction refers to the value of prediction of the mf model, while mf-dm refers to the effective system recovery after the decision-making takes place, i.e., real evaluation of mf prediction solution in the system. from t = 1 to t = 3 the mf prediction provides almost the same recovery efficiency as the hf (loss is below 0.1% in total travel cost), while at t = 4 there is a slightly larger loss of recovery cost when compared with the ga decision making supported by the hf model. the mf approach, however, consumes a small fraction of the decision time. in figure 4-(c) it is possible to infer that the ga optimal decision for recovery (team allocation) is achieved in mf for each day in less than 2 seconds, regardless of the ga parameters. such small decision-time has a large potential to increase the agility of this decision process. finding the optimal decision with ga relying on the hf technique would impose a delay of the decision-making scheme stalling the recovery process. decision-making that relies on the hf technique would require at least 25 minutes to achieve similar results in the most critical days, even considering that this is a relativity low dimensional problem in a relatively simple network. it is possible to infer that the analysis time increases significantly when the optimisation search in the ga uses more points, which is an important consideration for analyses that include more decision variables and that may demand larger population sizes. when the population size of 100 individuals is applied jointly with the convergence of the fitness function for 50 generations, the effectiveness of the hf dm making decreases substantially (decision times increase by more than 3); while in practical terms the mf model is largely unaffected by this. it is still capable of enabling zero-time responses in practice. this may be an important consideration in more involved applications. 4.2. adaptation of traffic network under incomplete information about the damage scenario the role of mf modelling in a context of adaptation is now discussed in the sioux-falls network, see figure 5. this network includes 76 links and 24 nodes. in the present example, 14 od pairs (with a demand of 200 users/hour for each od) are considered in the modelling, each comprising 4 routes. the capacity of each link is set initially to be 50% larger than the od pair demand. the idea of the current implementation of mf modelling is to find a representative example of the role of different fidelities in adaptation of a system. in this case decision-making for adaptation should be studied considering a scenario of incomplete information. a damage is known to have occurred in the network, however its extent is uncertain. existing information indicates that there is a likelihood of 70% that a scenario of minor damage occurred; where the number of affected links (n(ld)) has 25% probability of being n(ld) ≤ 4, 35% probability of being n(ld) = 5, 20% probability of being n(ld) = 6, 15% probability of being n(ld) = 7, and 5% probability of being n(ld) ≥ 8. in all possibilities the extent of the damage is expected to be 70% of the link capacity. then, there is a 30% probability of a scenario of high damage having occurred; where there is 30% probability of (n(ld)) ≤ 6, 20% probability of n(ld) = 7, a 20% probability of n(ld) = 8, a 15% probability of n(ld) = 9, and 15% probability of n(ld) ≥ 10. in all the cases of this scenario, a damage of capacity of 75% is expected. figure 5 presents potentially affected links (with highlighted link numbers). a decision should be made about how to adapt the network to this disruption. decisions that use the hf model are highly accurate, but involve large computational cost. in this example of disruption, with 18 dimensions, that would cover the decision to adapt potentially disrupted link capacities. so, as soon as the decision-maker commits to search for an optimal decision, he will need to perform an optimization 227 r. teixeira, b. martinez-pastor, m. nogal et al. acta polytechnica ctu proceedings figure 3. perturbation considered in the nguyen-dupuis network. red arrows represent the links that are damaged. figure 4. results of the application of the mf approach to the problem of recovery. (a) results based on a ga team distribution with a population size of 50 and that converge when no improvement in the fitness function occurs for 20 iterations. (b) results based on a ga team distribution that uses a population size of 50 and converges when no improvement in the fitness function occurs for 20 iterations. (c) computational times for (a)-(b). 228 vol. 36/2022 role of multi-fidelity in system adaptation and recovery figure 5. sioux-falls network and uncertain scenario of damage simulated. the (red) rectangles identify the links that are potentially damaged in the uncertain disruption. of the decision criteria on the hf model. there is limited adaptivity to this procedure, as changing the input conditions to it will involve a restart of the optimization calculations. if a mf model is introduced, then decision-making into adaptation can become adaptive; and because running an optimization scheme on a lf model that is representative of the hf model is virtually costfree, then, as new information becomes available it is possible to change decisions. in the case of the sioux-falls with 18 potentially damaged links, the lf model is built in slow-time and runs in tandem with the hf model considering an x that includes the capacities of each link from a reduction of 80% in relation to the initial capacity to an increase of 50% of the initial capacity. it uses 3205 evaluations of the hf model and is able to validate in average a random global sample within 10% of absolute error in prediction. if new points are required these can be added in slow-time or fast-time. figure 6 presents the results for adaptation to the uncertain scenario presented in this example. when the system is damaged it is assumed that a budget of adaptation is available that can add extra capacity to the network links. this budget has a limit 300 users/hour, and any of the potentially damaged links cannot have its capacity enhanced in more than 150 users/per hour. the adaptation decision-making problem involves then finding the optimal usage of this budget to improve the capacity of the potentially damaged links in figure 6, such that the loss in travel cost in the system is minimized. a genetic algorithm is used to allocate the additional capacity in the links considered in order to mitigate the effects of the perturbation. the population size is set to 100 figure 6. results for the adaptation probability distribution function using hf and mf modelling. individuals, complemented with convergence of the best fitness value for 50 iterations. loss of effectiveness for the adaptation decision is compared in terms of the reduction of travel cost that is possible to achieve with the it (i.e. distributing the budget available among links). the black dotted line presents the probability distribution of the response considering predictions using the mf model. because the mf model is virtually costless to run, the optimal operational point can be determined as new information about the extent of the damage enters the decision-making scheme, that is, an optimal adaptation strategy can be defined for all possible scenarios at virtually no cost (zero-response time). this curve is to the left of the remaining adaptation decisions that rely on hf modelling combined with the same decision technique. because these rely on the hf model it is not possible to change the input conditions of the decision-making scheme variables with the same level of flexibility. a single optimization to a specific scenario of damage will demand several hours to run. three scenarios are therefore tested; to optimize the adaptation of the network considering that all links are damaged at 70% of their capacity, optimize the adaptation of the network considering that all links are damaged at 75% of their capacity, and optimize the adaptation of the network considering that no links are damaged. in relation to the case of no adaptation, all contribute to an important improvement of the response; nonetheless, still outperformed by the adaptation with mf. the capability to change ad-hoc the decision with new information means total adaptive capacity. 5. conclusions the present work discussed the role that multi-fidelity modelling may have in the future of system analysis, with explicit relation to decision-making in recovery and adaption. while significant research has been directed to the need for accurate system models that can 229 r. teixeira, b. martinez-pastor, m. nogal et al. acta polytechnica ctu proceedings accurately inform decision-making, limited research has been directed to the practicality of these as a tool to inform decision schemes. while modelling of systems becomes more complex and accurate, so does the cost of it (effort required). in particular if paired with an optimization scheme, achieving an optimal decision can be quite resource and effort-demanding. in practice, and considering that systems modelling is inherently complex, this may not provide an efficient alternative for decision-making in scenarios of disruption. decisions in recovery and adaptation require zero-response times to be effective. as systems are dynamic, non-zero response times imply that at the moment a decision is achieved, there is a likelihood that it won’t be optimal or near-optimal; a consequence of the system’s evolution in time. a metamodel was used to construct a multi-fidelity model, where it fulfilled the role of low-fidelity model, that enables zero-response times in recover and adaptation of a system. the ideas of slow-time and fast-time were introduced and the effectiveness of multi-fidelity models emphasized in two examples of traffic networks. an adaptive system should include zero-time decision-making schemes, so that the optimal decision is enabled in each moment of operation, and any recovery or adaptation of the system is continuous with the system operation. different layers of fidelity have a key role in enabling this system behaviour. references [1] y. shang, m. nogal, h. wang, a. r. m. r. wolfert. systems thinking approach for improving maintenance management of discrete rail assets: a review and future perspectives. structure and infrastructure engineering 0(0):1–19, 2021. https://doi.org/10.1080/15732479.2021.1936569. [2] s. guo. https://towardsdatascience.com/an-intr oduction-to-surrogate-modeling-part-iii-beyond -basics-a60125767271. [3] a. i. forrester, a. sóbester, a. j. keane. multi-fidelity optimization via surrogate modelling. proceedings of the royal society a: mathematical, physical and engineering sciences 463(2088):3251–3269, 2007. https://doi.org/10.1098/rspa.2007.1900. [4] b. peherstorfer, k. willcox, m. gunzburger. survey of multifidelity methods in uncertainty propagation, inference, and optimization. siam review 60(3):550– 591, 2018. https://doi.org/10.1137/16m1082469. [5] m. g. fernández-godino, c. park, n.-h. kim, r. t. haftka. review of multi-fidelity models. arxiv preprint arxiv:160907196 2016. https://doi.org/10.2514/1.j057750. [6] q. zhou, y. wu, z. guo, et al. a generalized hierarchical co-kriging model for multi-fidelity data fusion. structural and multidisciplinary optimization 62:1885–1904, 2020. https://doi.org/10.1007/s00158-020-02583-7. [7] r. teixeira, m. nogal, a. o’connor. adaptive approaches in metamodel-based reliability analysis: a review. structural safety 89:102019, 2021. https://doi.org/10.1016/j.strusafe.2020.102019. [8] d. j. toal. some considerations regarding the use of multi-fidelity kriging in the construction of surrogate models. structural and multidisciplinary optimization 51(6):1223–1245, 2015. https://doi.org/10.1007/s00158-020-02583-7. [9] r. teixeira, b. martinez pastor, m. nogal, a. o’connor. metamodel-based metaheuristics in optimal responsive adaptation and recovery of traffic networks. sustainable and resilient infrastructure 2022. https://doi.org/10.1080/23789689.2022.2029325. [10] s. chakraborty, s. adhikari, r. ganguli. the role of surrogate models in the development of digital twins of dynamic systems. applied mathematical modelling 90:662–681, 2021. https://doi.org/10.1016/j.apm.2020.09.037. [11] r. teixeira, b. martinez-pastor, m. nogal, a. o’connor. reliability analysis using a multimetamodel complement-basis approach. reliability engineering & system safety 205:107248, 2021. https://doi.org/10.1016/j.ress.2020.107248. [12] m. nogal, a. o’connor, b. caulfield, b. martinezpastor. resilience of traffic networks: from perturbation to recovery via a dynamic restricted equilibrium model. reliability engineering & system safety 156:84–96, 2016. https://doi.org/10.1016/j.ress.2016.07.020. [13] b. martinez-pastor, m. nogal, a. o’connor, r. teixeira. transport network resilience: a mapping and sensitivity analysis strategy to improve the decision-making process during extreme weather events. international journal of critical infrastructures 17(4):330–352, 2021. https://doi.org/10.1504/ijcis.2021.120165. [14] r. teixeira, a. o’connor, m. nogal. probabilistic sensitivity analysis of offshore wind turbines using a transformed kullback-leibler divergence. structural safety 81:101860, 2019. https://doi.org/10.1016/j.strusafe.2019.03.007. 230 https://doi.org/10.1080/15732479.2021.1936569 https://towardsdatascience.com/an-introduction-to-surrogate-modeling-part-iii-beyond-basics-a60125767271 https://towardsdatascience.com/an-introduction-to-surrogate-modeling-part-iii-beyond-basics-a60125767271 https://towardsdatascience.com/an-introduction-to-surrogate-modeling-part-iii-beyond-basics-a60125767271 https://doi.org/10.1098/rspa.2007.1900 https://doi.org/10.1137/16m1082469 https://doi.org/10.2514/1.j057750 https://doi.org/10.1007/s00158-020-02583-7 https://doi.org/10.1016/j.strusafe.2020.102019 https://doi.org/10.1007/s00158-020-02583-7 https://doi.org/10.1080/23789689.2022.2029325 https://doi.org/10.1016/j.apm.2020.09.037 https://doi.org/10.1016/j.ress.2020.107248 https://doi.org/10.1016/j.ress.2016.07.020 https://doi.org/10.1504/ijcis.2021.120165 https://doi.org/10.1016/j.strusafe.2019.03.007 acta polytechnica ctu proceedings 36:224–230, 2022 1 introduction 2 multi-fidelity (mf) modelling 2.1 forms of multi-fidelity modelling 3 integration of mf modelling in adaptation and response of engineering systems 4 discussion on recovery and adaptation of a traffic network to a disruptive event 4.1 recovery of traffic network in a scenario of damage 4.2 adaptation of traffic network under incomplete information about the damage scenario 5 conclusions references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0005 acta polytechnica ctu proceedings 38:5–10, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague toward a critical assessment of indoor-environmental quality standards ardeshir mahdavia, ∗, christiane bergerb a department of building physics and building ecology, tu wien, karlsplatz 1040 vienna, austria b department of architecture, design and media technology, aalborg university, 9000 aalborg, denmark ∗ corresponding author: amahdavi@tuwien.ac.at abstract. building design and operation requirements regarding indoor-environmental quality (ieq) are of direct relevance to occupants’ needs. in most buildings, occupant-related requirements pertain to the provision of conditions that support optimal task performance and are subjectively perceived as comfortable or pleasant. standards and guidelines, which entail ieq requirements specifications are commonly viewed as the main sources of reference for practitioners, who are expected to follow the provisions in these documents and provide corresponding proof of compliance. however, actual code compliance processes are not always accompanied by critical reflections regarding the evidentiary basis of the included mandates and recommendations. it is thus necessary to critically scrutinise standards in view of explicit or implicit references to the scientific basis of the entailed mandates. the present contribution explores and details a path toward such a critical assessment. to this end, we considered several frequently referenced standards pertaining to thermal, visual, and air quality aspects of indoor environments. the results of this illustrative assessment effort highlight the scope and limitations of the standards’ default approaches to the definition of ieq-relevant requirements. moreover, they point to considerable gaps in the chain of evidence from standards’ immediate content to the underlying factual sources. keywords: indoor-environmental quality, standards, evidence. 1. introductory remarks for professionals and stakeholders involved in the building delivery process, codes, standards, and guidelines related to indoor-environmental quality (ieq) represent a critical source of information and guidance. these resources typically define and specify the spectrum of quality requirements (including those directly relevant to the needs of occupants) to be addressed in the course of buildings’ planning, erection, and operation. specifically, ieq-related requirements are directly relevant to occupants’ needs. broadly speaking, requirements pertaining to building occupants range from some rather basic principles (i.e., avoiding harm to human health, provision of the minimum thermal, visual, acoustic, and air quality conditions) to criteria that could be classified as less tangible (e.g., creating indoor environments that occupants perceive as being pleasing and pleasant). risks in the former category should be obviously avoided. those in the latter category could be presumably tolerated if the temporal extent of the exposure is limited. it is important to recognize that what constitutes a comfortable environment has a strong subjective background. but the same is not true of necessary conditions for occupants’ health. occupants do not always consciously perceive sources of adverse health effects. for instance, dangerous gases such as carbon monoxide and radon cannot be perceived by humans. standards represent a key source of information and support for professionals and stakeholders in the building delivery process. building designs and the actual buildings’ performance need to comply with the ieq-related mandates in these standards [1]. the question is, however, if by the virtue of their entailment in the standards, mandates and requirements can be automatically assumed to be entirely objective and based on solid scientific evidence. this, we argue, cannot be assumed a priori, but must be rather examined carefully. 2. in search of evidence one could assume that, in order to locate the scientific basis for standard-based ieq mandates, one would need to look no further than the evidence provided by disciplines such as physiology, psychology, and ergonomics. however, neither standards themselves, nor the related compliance processes appear to explicitly address arguments for the applicability of standards in general and the uncertainty of their recommendations in particular. this circumstance may have negative implications. professionals may concentrate solely on demonstrating compliance with minimum requirements instead of focusing on creating genuinely high-quality indoor environments. this underlines the importance of critical reflections regarding the availability and strength of sound scientific evidence behind the entailed prescriptions in the standards. in this context, a key question is as follows: do we 5 https://doi.org/10.14311/app.2022.38.0005 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en ardeshir mahdavi, christiane berger acta polytechnica ctu proceedings figure 1. illustrative exemplification of the chain of reference from certification systems, to different layers of standards, and technical literature. ([2–5], [6–10], [11], [12, 13]). sufficiently understand the influence mechanisms of indoor-environmental factors on the health, comfort, productivity, and well-being of the building occupants? despite the plethora of valuable research in the past, the identification and realization of occupants’ health and comfort requirements remains a formidable challenge. given the increasing popularity of various total building quality evaluation schemes and certificates, it would be useful, in principle, to examine not only commonly used ieq-related codes, standards, and guidelines (and their recommendations), but also building evaluation and certification systems in view of any cited theoretical background and empirical evidence. however, as illustrated in figure 1, common systems for building rating and certification (e.g., [14, 15]) appear not to entail explicit specifications of indoorenvironmental requirements. rather, they tend to refer to other thematically related national or international standards such as iso, en, din, or ansi [1]. some of these standards (e.g., din en 15251) refer, in turn, to other standards. there are very few cases, where standards include direct evidence and justification for the specified requirements. they do, however, include at times either general literature or specific studies (see “technical literature” box in figure 1). nonetheless, it is not a straight-forward matter to isolate, in the referenced technical literature, the evidence underlying the standards’ ieq-related criteria and the mandated values of the relevant occupantcentric performance indicators. 3. about formats for the specification of requirements for practitioners involved in the building delivery process, the primary utility of standards may be in the specification of ieq-relevant performance indicators and their respective mandated values. these values can be specified in terms of various formats. for instance, maximum allowable values may be mandated for parameters such as pollutants’ concentration, magnitude of glare, or sound level. likewise, minimum values can be specified, for instance, for illuminance levels or ventilation rates. the specification of performance indicator values can help organize and streamline the procedures aiming at demonstration of compliance with standards and general quality assurance. assuming that such values can be monitored after the commissioning phase, the examination of contractually specified performance can follow a rational path and liability issues can be resolved in an accountable man6 vol. 38/2022 toward a critical assessment of indoor-environmental quality standards figure 2. example of a standard-based ieq indicator (operative air temperature) and its value as a function of the outdoor air temperature (based on standards din en 15251 and din en iso 7730). figure 3. illustrative exemplification of standard-based specification of an ieq indicator (maximum permissible air flow speed) and its value as a function of indoor air temperature and turbulence intensity (based on standards din en 15251 and din en iso 7730). ner. requirements can be also formulated in terms of “zones”, involving the combined consideration of multiple variables and their values (e.g., thermal comfort zone that encompass air and radiant temperatures, air velocity, and humidity). to exemplify this, consider the following graphs (figures 2 and 3), which are based on standards din en 15251 and din en iso 7730. these graphs illustrate, in generic form, the dependency of the value of a performance indicator (i.e., recommended operative temperature, maximum allowable air flow speed) as a function of other variables (in this case, outdoor temperature, indoor air temperature, turbulence intensity). the logic behind these kinds of requirements in the ieq field has both similarities and dissimilarities with other code-based building regulations. basic building codes prescribe, for instance, minimum dimensions of certain architectural elements in building design and construction domains (e.g., width of doors, stairs, corridors). such requirements, for instance those pertaining to accessibility and universal design, are typically based on straightforward practical considerations such as human body and wheelchair dimension. they can be readily understood by designers and implemented in actuality. hence the compliance can be fairly easily examined in the course of building commissioning. but the translation of this approach, namely inference from scientific facts to design requirements, is in case of ieq-related requirements much more complex. it is thus important to explore if and to which extent ieq-related standards bolster their credibility by provision of explicit references to scientifically established evidence. 4. theory versus practice a cursory inspection of ieq-related standards leads to the conclusion that they entail much in terms of prescriptive requirements, but rather little in terms of explicit evidence behind those requirements. as mentioned before, rating systems typically refer to 7 ardeshir mahdavi, christiane berger acta polytechnica ctu proceedings figure 4. illustrative schema of indoor-environmental comfort equation components. thermal comfort visual discomfort presumed key indoor-environmental comfort parameters (independent parameters) ambient air temperature, radiant temperature, relative humidity, air velocity background luminance, glare source luminance indoor-environmental comfort constructs indicating building users’ subjective evaluation (dependent parameters) predicted mean vote, predicted percentage of dissatisfied unified glare rating (ugr), visual comfort probability (vcp) table 1. independent and dependent parameters with regard to thermal comfort and visual discomfort. standards, which sometimes refer to other standards. standards occasionally refer to technical papers and report, but a direct link from these resources to the stated requirements is rarely established. hence, such references appear to have the function of a general topical bibliography, rather than repositories of directly relevant evidentiary material. various reasons may be responsible for this circumstance. firstly, standards could be argued to represent resources for practitioners, rather than repositories of scientific knowledge. professionals involved in the building delivery process are more likely to consult standards in view of practical solutions and regulatory constraints, rather than contributors to their scientific understanding. secondly, ieq standards, in contrast to building construction and structure codes, cannot rely only on engineering sciences, but must also take into account findings from human and social sciences such as physiology, psychology, and sociology. this arguably adds an additional layer of complexity to the requirements specification process, given the significant role of subjective and qualitative parameters in the formation of occupants’ perception and evaluation phenomena. thirdly, the genesis of ieq standards does not follow in all cases an entirely systematic process. thereby, factors other than the occupants’ health or comfort requirements may play a role. these other factors could include financial issues, policy considerations, or special interests. as with many other areas involving regulatory aspects, development of standards may require, from the involved stakeholders (e.g., government, industry, academia), various degrees of consent and the willingness to comprise. consequently, not all aspect of the resulting standards could be expected to strictly adhere to scientific evidence. the above reflections, however, should not be misunderstood. many commonly applied ieq standards do not explicitly elaborate on their underlying evidentiary basis. but they do provide implicit pointers to the kinds of principles and methods that inform them. in other words, standards frequently entail features that point a kind of implicit methodological and theoretical underpinnings. it may be thus interesting and useful, to engage in a kind of reverse engineering of standards and their syntax, terminology, and logic, so as to uncover their implicit theoretical foundations. 5. sources of indoor-environmental comfort constructs conventionally, indoor-environmental comfort equations are used in standards focusing on thermal and visual comfort. figure 4 schematically illustrates the elements of such indoor-environmental comfort equations. thereby, a set of physical variables is intended to formulate dependent variables that capture building users’ level of comfort [1]. whereas physical variables can be typically derived by measurements, dependent variables result from subjective evaluations. table 1 shows such constructs that are assumed to capture occupants’ subjective evaluations concerning thermal and visual comfort/discomfort. thermal comfort standards typically refer to relevant (measureable) physical variables, such as air temperature, air speed, and water vapour concentration. moreover, thermal comfort constructs capture subjective building users’ evaluations by means of scales that are commonly used in studies in psychology. the underlying logic to map operations in the comfort equation often relies on two sources. on the one hand, the source is based on underlying physiological mechanisms. for instance, the humans’ thermo-regulatory system to keep the human body’s kernel temperature is of key significance when addressing thermal comfort [16]. on the other hand, occupants’ evaluations 8 vol. 38/2022 toward a critical assessment of indoor-environmental quality standards of experiments with different indoor-environmental conditions provide another source of information. in which way these two sources contribute in order to derive a comfort equation may be quite different. concerning thermal comfort, the understanding of thermo-regulatory processes is crucial in order to formulate thermal comfort models. moreover, studies involving subjective evaluations of participants provide relevant data of physiological-relevant variables. with regard to visual comfort, a physiological understanding of glare or light scattering is essential. discomfortable visual indoor-environmental conditions are primarily evaluated by occupants’ subjective evaluations. 6. reflections on limitations codes, guidelines, and standards seldom bring to light the relevant and precise explanation for the particular requirements they demand, some of which provide exact limitations for or allowed ranges of different variables. previously, we suggested some reasons for this situation. these include difficulties in accurately formulating occupant-oriented concepts for comfort, well-being, and health; recognising and estimating the proper ieq proxies; the multi-domain character of indoor-environmental circumstances; the variance of occupants’ requirements and the constantly changing nature of these requirements; the actual challenges of working procedures in standardization bodies. it is important to emphasize that we did not embark on the present treatment with an unreasonably high expectation. clearly, it is not likely that a single formula or scheme could provide definitive evidence for all certification systems and regulatory guidelines regarding indoor environment. however, this process revealed a repeated pattern, known from the fields of psychology, physiology, and medicine, which can be described in the following way. in standards, the values and ranges of the relevant variables meant to be representative of the indoor environment conditions are mapped onto the values of specific indicators of human comfort and health. this mapping process usually relies on the combination of two components, specifically (1.) a psychologically or physiologically based theory, and (2.) obtainable research data from experiments with human participants. in cases where explicit comfort calculations are needed (e.g., visual discomfort or thermal comfort), the guidelines provide specific models defining occupant comfort and health (including their required values and ranges). otherwise, these models may not be specifically provided. rather, the assumption appears to be that when specific indoor-environmental variables (e.g., indoor air carbon dioxide concentration) are kept in mandated ranges, the requirements related to occupants’ comfort and health are met. we can summarize the foregoing discourse regarding the current and future state of guidelines, codes, standards as follows: • to begin with, we need to acknowledge that it is at times difficult to follow the trace from guidelines back to their evidentiary basis. in practice, this may lead to a situation, where standards and their requirements do not serve as reliable sources of disciplinary guidance, but are reduced to opaque propositions that are thoughtlessly followed. it is not our intent to suggest that standards should reproduce the complete theoretical substructure and scientific evidence that they are built upon. their focus needs to be on operational matters and specific directives, as they mainly serve as regulative tools. however, it would be justified to look for an explicit link between the standard-based requirements and their scientific basis. this link appears to be frequently untraceable. • secondly, we face numerous challenges, while trying to identify the theories and supporting data that would document the credibility of the requirements formulated the standards. research in the fields of human health, well-being, and comfort continuously yields new and valuable results, but it also involves numerous challenges and uncertainties. while reputable scientists in the field commonly avoid expressing doctrinal standpoints and claiming the absolute truth, regulating authorities are obligated to derive from insufficiently uncertain knowledge mandatory rules and regulations. as such, it seems as if a certain level of uncertainty in the objectivity of mandated requirements is the price one is willing to pay in order to prevent chaos, confusion, and accountability issues in the building design and delivery process. • thirdly, the identified ieq-relevant standardization challenges (especially, the scantiness of explicit arguments for the entailed requirements) may also point to gaps in our scientific understanding in the related domains. it is not always clear which physical characteristics of the indoor environment represent the “correct” variables to be considered when evaluating comfort and health. further difficulties must be encountered when we try to define and validate the constructs for comfort and health that do not fully capture the relationship between psychological, physiological, and social aspects of occupants’ perception and assessment of indoor environment. moreover, the scope of available empirical data for the evaluation and validation of behavioural and perceptual theories is still rather limited. the indoor-environmental exposure circumstances are fundamentally multi-domain. hence, isolated treatment of these separate ieq domains is likely to fail in capturing their complexity. we have made some progress in our appreciation of occupants’ diversity and the dynamic nature of their preferences. 9 ardeshir mahdavi, christiane berger acta polytechnica ctu proceedings however, probably the level of progress has not been high enough to warrant a translation of the respective findings in the appliable ieq regulations. the investigation of the formulation and implementation of ieq-relevant standardization procedures structures points not only to their limitations, but also to certain shortcomings in our scientific understanding of occupants’ preferences and necessities in indoor environments. there is a need for clearer, better structured and more objective procedure, when it comes to implementing our current scientific knowledge, with all its undeniable limitations, into ieq standards, guidelines, and codes. at the same time, we urgently need to extend and refine the depth of our knowledge regarding the mechanisms by which indoor-environmental conditions influence human health, comfort, wellbeing, and productivity. references [1] a. mahdavi, c. berger, v. bochukova, et al. necessary conditions for multi-domain indoor environmental quality standards. sustainability 12(20):8439, 2020. https://doi.org/10.3390/su12208439. 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[2021-10-14], https://www.usgbc.org/leed. [15] dgnb system – sustainable and green buildings. [2021-10-14], https://www.dgnb-system.de/en/. [16] a. mahdavi. the human factor in sustainable architecture. in k. al-sallal (ed.), low energy low carbon architecture: recent advances & future directions, pp. 177–198. 2017. https://doi.org/10.1201/b19882-12. 10 https://doi.org/10.3390/su12208439 https://doi.org/10.1111/j.1600-0668.2004.00270.x https://doi.org/10.1016/s0378-7788(02)00013-0 https://doi.org/10.1016/0360-1323(94)e0006-d https://www.usgbc.org/leed https://www.dgnb-system.de/en/ https://doi.org/10.1201/b19882-12 acta polytechnica ctu proceedings 38:5–10, 2022 1 introductory remarks 2 in search of evidence 3 about formats for the specification of requirements 4 theory versus practice 5 sources of indoor-environmental comfort constructs 6 reflections on limitations references my title https://doi.org/10.14311/app.2022.36.0280 acta polytechnica ctu proceedings 36:280–285, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague effect of decision parameter efficiency on target reliability andrew way∗, nico de koker, celeste viljoen stellenbosch university, faculty of civil engineering, hammanshand road, stellenbosch, 7600, south africa ∗ corresponding author: acway@sun.ac.za abstract. target reliability forms the basis of most modern design standards and is intended to represent an optimal balance of safety and risk of failure. previous research noted discrepancies between target reliability obtained using generic and case-specific cost optimization for sls cases. the source of the discrepancies was identified as the efficiency of the decision parameter at increasing reliability. this research extends the investigation to cases of uls failures and found similar discrepancies. generic cost optimization assumes a linear dependence between the decision parameter and the structural resistance. where the dependence of resistance on the decision parameter is superlinear, the generic was found to under-predict target reliability by up to 15%. a factor is proposed to adjust generically-obtained uls target reliability to be more appropriate to specific uls cases. the factor accounts for the efficiency of the decision parameter, the case-specific cost of safety and parameter variation. the adjustment factor represents a first step towards mapping generic to case-specific target reliability in the ideal of promoting safer and more cost-effective structures. keywords: cost optimization, structural reliability, target reliability. 1. introduction target reliability in structures is determined through the optimization of costs, considering the consequences of failure [1]. where human safety is at risk in the case of failure, lower limits on reliability are imposed based on societal risk criterion [2]. the costs of increasing safety and the consequences of failure play an important role in the determination of target (optimum) reliability and have been considered extensively in various research contributions [2–6]. the effect that the decision parameter efficiency has on the reliability limit state, however, has not been given as much consideration. in the cost optimization process, a function describing the structural cost is developed, in conjunction with the reliability limit state that governs the design of a structure. the parameter with the greatest influence on increasing the reliability in the limit state is identified as the decision parameter. the optimal reliability is found at the decision parameter value that minimizes the structural cost function. previous research [7–9] has shown that in rc structures governed by sls design, the efficiency of the decision parameter at increasing safety in the limit state equation also plays a role in the determination of the optimum reliability. generic target reliability through cost optimization is typically based on research by rackwitz [4]. in this generic assessment, the structural resistance is linearly dependent on the decision parameter. however, when the decision parameter in the governing limit state takes on a form that is not linear, the target reliability will differ from that obtained through the generic cost optimization. this research aims to investigate various uls limit states to determine the extent to which the decision parameter efficiency affects target reliability and how deviations from the generic target reliability may be predicted and accounted for. 2. target reliability from cost optimization 2.1. generic cost optimization a summary of the generic cost-optimization framework to determine target reliability is presented here, from [4]. the total cost of the structure normalized by the initial structural cost, z(d), is shown in equation 1 as a function of a decision parameter, d. z(d) = cp c0 + u c0 · pf sls γ + ( cp c0 + a c0 ) · ω γ + ( cp c0 + h c0 ) · pf u ls γ (1) with cp c0 + c1 · d c0 initial cost of structure independent of d c1 initial costs of structure dependent on d γ age-averaged discount rate ω obsolescence rate u, a, h costs related to sls failure, obsolescence and uls failure pf probability of failure costs attributed to inspection and maintenance are assumed to be included in cp; costs related to fatigue and ageing failure are not considered. the reliability 280 https://doi.org/10.14311/app.2022.36.0280 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 36/2022 effect of decision parameter efficiency on target reliability limit states and probabilities of failure for uls and sls that are linked to the cost function in eq. 1 are given as: gu ls = ru ls (d) − s (2) pf,u ls = p (gu ls < 0) (3) gsls = rsls (d) − s (4) pf,sls = p (gsls < 0) (5) the decision parameter that which most efficiently increases reliability at a specific limit state. in a generic case, it typically affects only the structural resistance, which is modelled lognormally, with mean and coefficient of variation (vi) as ru ls ∼ ln (d, 0.3) and rsls ∼ ln (d/1.5, 0.3) for uls and sls, respectively [4]. in effect, this definition reflects the fact that the loading at uls is typically ≈ 1.5 times that at sls and that the decision parameter required at sls is assumed to be affected in a similar fashion. the load is therefore modelled as s ∼ ln (1, 0.3) for both uls and sls. target reliability from cost optimization is defined by the value of decision parameter which minimizes the total cost of the structure. this corresponds to the point d : ∂z(d)/∂d = 0. from eq. 6, the optimal value of d is dependent on various cost ratios, as well as γ and ω, all of which have been investigated in previous research [3, 10–13]. ∂z(d) ∂d = 0 = ( 1 + h c0 + c1d c0 ) ∂pf,u ls ∂d + c1 c0 ( ω + 1 γ + 1 ) pf,u ls + u c0γ ∂pf,sls ∂d (6) little consideration has been given, however, to the derivative of the probability of failure with respect to the decision parameter at uls and sls (∂pf,u ls /∂d and ∂pf,sls /∂d, respectively). this "efficiency parameter" describes how the probability of failure is affected by changes in d and is a measure of how efficient d is at increasing the reliability of the limit state. 2.2. efficiency parameter from eq. 2 to 5, the efficiency of the decision parameter at increasing safety is dependent on the limit state equations for uls and sls. previous research by van nierop et al. [7–9] investigated target reliability in reinforced concrete (rc) structures where the design is governed by sls considerations. this work indicated that discrepancies exist between the target reliability derived using the generic cost optimization and that derived from a case-specific cost optimization. this discrepancy was attributed to the difference in efficiency of the decision parameter in the generic sls formulation compared to that of the specific case, i.e. generic︷ ︸︸ ︷ ∂pf,sls ∂d ̸= specif ic︷ ︸︸ ︷ ∂pf,sls ∂d and proposed a parameter, ν, that could be used to map the specific case to the generic, of the general form: generic︷ ︸︸ ︷ ∂pf,sls ∂d = ν · specif ic︷ ︸︸ ︷ ∂pf,sls ∂d . in the context of target reliability, however, it is more convenient to consider it in terms of the reliability index, as the conversion from probability of failure to reliability index, β, is case-independent: generic︷ ︸︸ ︷ ∂βgen ∂d = ν · specif ic︷ ︸︸ ︷ ∂βspec ∂d . (7) van nierop et al.[8] investigated this discrepancy in target reliability for rc structures governed by sls design. in the current research, the efficiency parameter is investigated in a broader structural context to determine the extent to which it affects target reliability in general. 3. effect of decision parameter on target reliability 3.1. form of decision parameter in limit state equations to investigate the effect of the decision parameter efficiency on target reliability, the governing limit state equation, g (eq. 2 5), is considered in more detail. in the generic cost optimization, a linear dependence is assumed between the decision parameter and the resistance in the generic limit state equation. that is, linear increases in the decision parameter linearly increase the mean of the resistance, and thereby increase the reliability. a typical example of this is the bending resistance of a simply supported rc beam, given by the simplified limit state in eq. 8, where as, fy , x, w and l are the area of tension reinforcement, reinforcement yield stress, moment lever arm, distributed load and length, respectively. the decision parameter is as, which clearly has a linear effect on the resistance and therefore the limit state. in cases like this, the assumption of linearity, and by implication the target reliability derived from it, will also be appropriate. g1(as) = fy · as · x − w · l2 8 (8) in cases where the decision parameter form in the limit state equation is not linear, whether sub-linear or super-linear, the generic assumption of linearity is no longer appropriate. consider the tension resistance of a structural steel rod with radius r, subjected to a tensile load, tu, as shown in the limit state in eq. 9. 281 a. way, n. de koker, c. viljoen acta polytechnica ctu proceedings comparing eq. 8 to eq. 9, it is clear that increasing r is more efficient at increasing the resistance than increasing as, and thus that ∂βg1 /∂as < ∂βg2 /∂r. in the latter case, a higher target reliability will result. g2(r) = πr2fy − tu (9) this suggests that it is the form of the dependence of the governing limit state equation on the decision parameter that governs the efficiency of the decision parameter and thereby the target reliability. decision parameter efficiency will apply to any structure and not just rc structures governed by sls requirements. cases of uls limit states are investigated below. 3.2. uls test cases the target reliability of a number of specific uls limit states are determined and compared with those obtained from generic cost optimization. varying degrees of decision parameter efficiency are implied by the chosen limit state equations, to determine the effect that the decision parameter form has on the target reliability. a more generalized structural cost equation is used to determine the target reliability as shown in eq. 10, where cf are all costs associated with failure of the limit states under consideration. the normalized failure cost is varied, with particular points of interest at cf /c0 = 1, 3.5 and 7.5, which are representative of small, medium and large consequences of failure (cof), respectively [11]. similarly, the normalized costs of increasing safety (cos) are varied, using c1/c0 = 5 × 10−2, 5 × 10−3, and 5 × 10−4 for high, moderate and low costs of increasing safety, respectively [3, 14]. furthermore, dn in eq. 10 represents a normalized decision parameter, as discussed below. z(d) =1 + c1dn c0 + ( 1 + c1dn c0 + a c0 ) ω γ + ( cf c0 ) pf γ (10) for the intents of evaluating the effect of decision parameter efficiency on target reliability, the rest of the cost optimization parameters are taken directly from [4], with the exception of the case of lower variation, and are shown in table 1. the various uls failure mechanisms in table 1 are chosen to illustrate the effect of decision parameter efficiency on target reliability, in comparison to that obtained from the generic cost optimization. the following assumptions/simplifications are made, and are discussed below: • all action effects (s, m, t, n ) are considered with both high and low vs and are modelled as ln (µs = 1, vs = 0.3) and ln (1, 0.15), in turn; • all resistance effects are modelled with a mean value stemming from the resistance term from table 1 and vr = 0.3 and 0.1, in turn. e.g. for case 2: r ∼ ln (fy as z, 0.3) and ln (fy as z, 0.15), in turn; • all decision parameters are normalized with respect to the value of decision parameter, d0, that gives β = 0, i.e. dn = d/d0; • model factors are included as part of vi. the above-mentioned assumptions are made so that meaningful comparisons can be drawn between decision parameter efficiencies, in terms of target reliability. in reality, the statistical distribution and variation of each resistance and load parameter in the various limit states will vary; they are kept consistent here, for the sake of comparison with the generic case. similarly, the model factors assumed to be included in vi will vary for each case. the normalization of the decision parameter is performed so that the normalized costs of increasing safety (c1/c0) can be used for all the cases in table 1. this avoids the subjectivity of considering costs specifically for each case of decision parameter, and is similar in principle to that performed for target reliability in existing structures from [15]. all reliability analyses are performed using the first order reliability method. 4. results and discussion the results of the optimization for higher variation are shown in 1 and 2. red vertical lines indicate small, medium and large cof. the magnitudes of the target reliability in figures 1 and 2 is lower than typical due to the high variation, which is intended to cater to a wide array of applications. from figures 1 and 2, the case of concrete column buckling (decision parameter of a 4th order) has the highest target reliability. uls case limit state equation decision parameter form (order) 1 generic r(d) − s generic decision parameter d linear (1st) 2 rc beam bending fy · as · z − m area of tension reinforcing as linear (1st) 3 steel tension rod πr2fy − t steel rod radius r quadratic (2nd) 4 rc column buckling ecπ3d4/64l2 − n concrete column diameter d quartic (4th) cost optimization parameters γ = 0.035 ω = 0.02 a/c0 = 0.2 cf /c0 : 0.1 − 7.5 c1/c0 : 5 × 10−2; 5 × 10−3; 5 × 10−4 table 1. uls test case limit state equations and cost optimization parameters. 282 vol. 36/2022 effect of decision parameter efficiency on target reliability 0 1 2 3 4 5 6 7 cf/c0 10 6 10 5 10 4 10 3 p f ,1 vr, vs = 0.3, 0.3 generic rc bending tie rod column buckle 0 1 2 3 4 5 6 7 cf/c0 3.00 3.25 3.50 3.75 4.00 4.25 4.50 4.75 5.00 t, 1 vr, vs = 0.3, 0.3 moderate cos generic rc bending tie rod column buckle figure 1. annual probability of failure (left) and annual target reliability (right) for the considered uls limit states for moderate cos (c1/c0 = 5 × 10−3) with high variation. the difference in target reliability with respect to the generic case is ∆βt,1 ≈ 0.6 for all three cases of cos. the generic case and that of bending in an rc beam (both linear decision parameters) have the lowest target reliability and are practically identical. the case of a steel tension rod (quadratic decision parameter) falls in between the generic and concrete buckling cases, with ∆βt,1 ≈ 0.3. these results confirm that target reliability increases with increasing decision parameter efficiency. the process is repeated for lower values of vr = 0.1 and vs = 0.15 to determine whether or not similar trends apply to limit states with variation more representative of typical applications [2, 16], where parameters and/or model factors exhibit lower variation. for the sake of brevity, only the results for moderate cos are shown. from figure 3, similar trends can be observed for lower variation, with higher values of target reliability. similar trends were observed for cases of low and high cos (not shown here). differences in magnitude of target reliability between the generic and column buckling cases are slightly lower for lower variation, at ∆βt,1 ≈ 0.5. in the consideration of target reliability, higher values are justified when the cof are large, when parameter and model uncertainty is low or when the costs of increasing safety are low. in this case, the latter is true for uls cases 3 and 4. as a result of the heightened decision parameter efficiency, the cost of increasing safety is effectively lowered. the generic cost optimization under-predicts the target reliability for these cases due to the assumed linear form of the relationship between the decision parameter and the resistance term in the limit state equation. figure 4 illustrates this visually by comparing the target reliability obtained from the generic with that from the specific, for moderate cos, high variation and a range of cof of 0.1 ≤ cf /c0 ≤ 7.5. similar results were found for other cos and values of variation. thus, for any specific limit state with a decision parameter form other than linear, the target reliability obtained from the generic cost optimization will need to be adjusted using a factor that considers the decision parameter efficiency. 0 1 2 3 4 5 6 7 cf/c0 3.75 4.00 4.25 4.50 4.75 5.00 5.25 t, 1 vr, vs = 0.3, 0.3 low cos generic rc bending tie rod column buckle 0 1 2 3 4 5 6 7 cf/c0 2.25 2.50 2.75 3.00 3.25 3.50 3.75 4.00 4.25 4.50 t, 1 vr, vs = 0.3, 0.3 high cos generic rc bending tie rod column buckle figure 2. annual target reliability for the considered uls limit states for low and high cos (c1/c0 = 5 × 10−4 and 5 × 10−2, from left to right, respectively) with high variation. 283 a. way, n. de koker, c. viljoen acta polytechnica ctu proceedings cost of safety vi low moderate high form of decision parameter in governing limit state linear quadratic quartic linear quadratic quartic linear quadratic quartic high 0.0 0.3 0.6 0.0 0.3 0.6 0.0 0.4 0.7 low 0.0 0.2 0.4 0.0 0.2 0.4 0.0 0.4 0.7 table 2. target reliability adjustment factor (ν) values for various forms of decision parameter, with reference to eq. 11 for all considered consequence of failure classes (0.1 ≤ cf /c0 ≤ 7.5). 0 1 2 3 4 5 6 7 cf/c0 3.6 3.8 4.0 4.2 4.4 4.6 4.8 5.0 5.2 t, 1 vr, vs = 0.1, 0.15 moderate cos generic rc bending tie rod column buckle figure 3. annual target reliability for the considered uls limit states for moderate cos (c1/c0 = 5 × 10−3) with low variation. from the results, an additive factor of the form in eq. 11 captures the deviation from the generic target reliability more effectively than the multiplicative form proposed in [7]. generic︷ ︸︸ ︷ βt,1,gen + ν = specif ic︷ ︸︸ ︷ β1,t,spec (11) values for the target reliability adjustment factor, ν, are shown in table 2 for first, second and fourth order decision parameter forms. these values are applicable for low, medium and high cof and are given for high (vr, vs = 0.3, 0.3) and low (vr, vs = 0.1, 0.15) variation. it should be noted however, that these are indicative first estimates and that further research is required, especially for cases where the cos of a specific decision parameter varies substantially from the considered ratios. additionally, the determination of similar values for sls cases should be considered in future research. 5. conclusions target reliability determined by generic cost optimization forms the basis of design of most modern design codes. to maintain acceptable levels of reliability and simultaneously design cost effective structures, target reliability should be as reflective of the specific structure under consideration as possible. previous 3.25 3.50 3.75 4.00 4.25 4.50 4.75 gen 3.25 3.50 3.75 4.00 4.25 4.50 4.75 sp ec vr, vs = 0.3, 0.3 moderate cos generic rc bending tie rod column buckle figure 4. comparison of annual target reliability for generic vs specific cases for moderate cos, high variation and 0.1 ≤ cf /c0 ≤ 7.5 research reported discrepancies between sls target reliability obtained using generic cost optimization and that obtained using structure-specific cost optimization. the efficiency of the decision parameter at increasing safety was identified as the cause of the discrepancies. this research investigated cases of uls failures and also identified discrepancies in target reliability of up to ∆β = 0.7 between generic and specific optimization. for forms of decision parameter in the limit state equation other than linear, the target reliability differs from that obtained through generic cost optimization and needs to be adjusted to be representative of the specific case. a factor is sproposed to appropriately adjust the target reliability, based on decision parameter efficiency, parameter variation and cost of safety. future research aims to determine similar values for sls, as well as a more comprehensive consideration of decision parameter forms in governing limit states. acknowledgements this research was supported by post-doctoral funding from research subcommittee b of stellenbosch university. references [1] iso. iso 2394:2015 general principles on reliability for structures. international organization for standardization (iso), geneva, switzerland, 2015. 284 vol. 36/2022 effect of decision parameter efficiency on target reliability [2] r. caspeele, r. steenbergen, m. sykora. partial factor methods for existing concrete structures. tech. rep., fib bulletin 80, 2016. [3] k. fischer, c. viljoen, j. köhler, m. faber. optimal and acceptable reliabilities for structural design. struct safe 76:149–161, 2019. https://doi.org/10.1016/j.strusafe.2018.09.002. [4] r. rackwitz. optimization the basis of codemaking and reliability verification. struct safe 22:27–60, 2000. https://doi.org/10.1016/s0167-4730(99)00037-5. [5] o. ditlevsen, h. madsen. structural reliability methods, vol. 178. new york: wiley., 1996. [6] m. holický, d. diamantidis, m. sýkora. determination of target safety for structures. in 12th international conference on applications of statistics and probability, vancouver. 2015. https://doi.org/10.14288/1.0076258. [7] s. van nierop. target reliability of concrete structures governed by serviceability limit state design. master’s thesis, stellenbosch university, stellenbosch, south africa, 2017. [8] s. van nierop, c. viljoen, r. lenner. target reliability of concrete structures governed by serviceability limit state design. 15th international probabilistic workshop, dresden, germany 2017. [9] r. lenner, c. viljoen, s. van nierop. a comparative study of target reliability index derivation for reinforced concrete structures governed by serviceability limit state. struct conc 20(2):670–677, 2019. https://doi.org/10.1002/suco.201800202. [10] d. diamantidis, m. sýkora, e. bertacca. obsolescence rate: framework, analysis and influence on risk acceptance criteria. in life cycle analysis and assessment in civil engineering: towards an integrated vision: proceedings of the sixth international symposium on life-cycle civil engineering (ialcce 2018), ghent, belgium. crc press., vol. 5, p. 379. 2018. [11] joint committee on structural safety (jcss). probabilistic model code, 2001. [12] r. rackwitz. optimization with lqi acceptance criterion. munich, germany: jcss basic documents on risk assessment in engineering, 2008. [13] r. van coile, d. hopkin, l. bisby, r. caspeele. the meaning of beta: background and applicability of the target reliability index for normal conditions to structural fire engineering. in 6th international workshop on performance, protection & strengthening of structures under extreme loading, protect2017, guangzhou (canton), china. 2017. https://doi.org/10.1016/j.proeng.2017.11.110. [14] k. fischer, b.-v. c., f. m.h. deriving target reliabilities from the lqi. in lqi symposium, kongens lyngby, denmark. 2012. [15] m. sýkora, d. diamantidis, m. holický, k. jung. target reliability for existing structures considering economic and societal aspects. structure and infrastructure engineering 13(1):181–194, 2017. https://doi.org/10.1080/15732479.2016.1198394. [16] m. holický. reliability analysis for structural design. sun press, sun media, stellenbosch, south africa., 2009. https://doi.org/10.18820/9781920689346. 285 https://doi.org/10.1016/j.strusafe.2018.09.002 https://doi.org/10.1016/s0167-4730(99)00037-5 https://doi.org/10.14288/1.0076258 https://doi.org/10.1002/suco.201800202 https://doi.org/10.1016/j.proeng.2017.11.110 https://doi.org/10.1080/15732479.2016.1198394 https://doi.org/10.18820/9781920689346 acta polytechnica ctu proceedings 36:280–285, 2022 1 introduction 2 target reliability from cost optimization 2.1 generic cost optimization 2.2 efficiency parameter 3 effect of decision parameter on target reliability 3.1 form of decision parameter in limit state equations 3.2 uls test cases 4 results and discussion 5 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0281 acta polytechnica ctu proceedings 38:281–287, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague building-integrated photovoltaics (bipv) combined with hydrogen-based electricity storage system at building-scale towards carbon neutrality sergi aguacila, ∗, yvan morierb, philippe coutyb, c, jean-philippe bacherb a ecole polytechnique fédérale de lausanne (epfl), building2050 group, fribourg, switzerland b university of applied science of western switzerland (heia-fr, hes-so), energy institute, fribourg, switzerland c tecphy sàrl, renens (vd), switzerland ∗ corresponding author: sergi.aguacil@epfl.ch abstract. electricity storage technologies in buildings are evolving, mainly to reduce their environmental impact and to improve self-sufficiency of buildings that produce their own energy through building-integrated photovoltaics (bipv) installations. to maximize self-consumption – minimizing the import of grid electricity – photovoltaic (pv) systems can be coupled with a hydrogen storage system converting the electricity to hydrogen by electrolysis during the summer season – when the on-site production is higher – and employing it during the winter season with fuel cells. this study focuses on the sizing constraints of solar hydrogen systems at building-scale using an innovative research-centre that will be built in fribourg (switzerland). it presents four stories and a mix-usage (office spaces and research facilities areas) and multi-oriented pv installation in order to produce enough electricity to achieve at least 50 % of electricity self-sufficiency ratio. using the pv production, this study aims to optimise the sizing of a hydrogen storage system allowing to reach the required self-sufficiency ratio with the lowest environmental impact possible. ultimately, the global energy and financial efficiency of the system will be analysed. keywords: building-integrated photovoltaics, building energy analysis, solar hydrogen storage, fuel cell, battery storage system. 1. introduction photovoltaic offers an easy and affordable manner to produce sustainable electricity using solar power. over the last decade, the number of solar panels installation in switzerland has been multiplied by five [1] and the overall electricity efficiency of these panels has kept increasing. solar energy will play a big role in the path to reach carbon neutrality by 2050. however, one issue remains unchanged, solar energy is only available during a short period of time during the day and in most countries, the production varies greatly throughout the year. this means that great selfsufficiency (ss) ratios (equation (1)) can be achieved during the day and summer but energy needs to be stored in order to be reused during the night or winter when the production isn’t sufficient. ssratio = p v production used onsite t otal electrical consumption (1) daily storage is usually achieved by using battery packs, however, they have a large carbon footprint [2]. nowadays, in most pv installations, very little storage is integrated allowing the building to be self-sufficient for 4 to 6 hours [3, 4]. this paper focuses on the use of seasonal hydrogen (h2) storage at buildingscale. in the summer, excess of production can be used to produce h2 by electrolysis of water and can then be stored in individual tanks. in the winter, when the electrical demand is higher than the on-site production, the h2 can be used in a fuel cell to produce electricity. this process also releases heat which can be recovered for heating or domestic hot water (dhw) converting the fuel cell in a co-generation unit. to simulate the h2 cycle over the year, a ms excel tool has been developed in order to simulate the way the energy is stored and reused. the program allows for two types of storage, a li-ion battery pack and a hydrogen storage system. in order to run a simulation, the program requires the following main input data: • hourly time-step electricity pv production and electrical and thermal demand [kwh]. • h2 storage specification (power and efficiency of the fuel cell and electrolyser, storage capacity). • battery pack specification (capacity and efficiency). • thermal storage specification (hot water tank volume, tank’s insulation, storage temperature). • additional data on components (life cycle, cost, greenhouse gases (ghg) emission). using this information, the program simulates the 281 https://doi.org/10.14311/app.2022.38.0281 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en s. aguacil, y. morier, p. couty, j.-p. bacher acta polytechnica ctu proceedings energy storage during a year and can display output values such as ss ratio, amount of electrical energy imported and exported from or to the grid, amount of thermal energy stored, the overall efficiency of the h2 cycle or the amount of ghg emitted. in a first scenario, the sizing of a h2 storage system allowing to reach 100 % ss was computed. this showed that to achieve such a goal, the amount of h2 stored had to be enormous, around 75 days of autonomy and 1.4 tons of h2. in this paper, the goal is to reach an ss ratio of 80 %. moreover, without using heat recovery on the fuel cell, the overall efficiency of the h2 storage system only reached about 30 %. in a further study [5], the focus was placed on heat recovery on the fuel cell. this time, parts of the simulation were made using polysun software [6]. heat recovery allowed the system to reach 43 % overall efficiency; however, limitations were quickly encountered. using the polysun software, we had very little control of the fuel cell as the program would only consider the thermal demand to switch it on. this is why this article focuses on the writing of an in-house excel tool that allows the user to control every energy flux (electrical and thermal). the global efficiency of the system can be optimized by controlling the electrolyser and the fuel cell in the most efficient way using different control modes that are presented in this paper. an effective sizing of the h2 storage system is then possible and outputs values are available to analyse the project. in order to run the simulation, this paper uses the smart living lab (sll) [7] as a real case study. the building will already be equipped with energy storage in the form of li-ion batteries but will provide enough space for further research, in our case, a h2 storage system with a co-generation fuel cell allowing to produce heat and electricity for the building. 2. methodology the methodology encompasses three phases: (1.) hourly energy demand and production are estimated using the energy modelling and analysis of the building. the thermal and electrical demands are computed using designbuilder [8] with energyplus simulation engine [9]. the pv production is calculated using the polysun software [6]. these data will be used in the further phases in the form of hourly time-step value. (2.) the implementation of a hydrogen-based storage system. in order to simulate the addition of a h2 storage system a dedicated program was written using a combination of ms excel and vba coding. this program includes a li-ion battery pack mainly for daily energy demand and a h2 system for seasonal storage. heat recovery on the fuel cell, hot water storage, and heat distribution are also included in the computation. (3.) the sizing and optimization process. the sizing process is made with the help of design explorer [10]. the optimization process then consists of tweaking the program in order to obtain different improved results such as greatest ss ratio, lowest ghg emission or highest overall efficiency for the h2 system. 2.1. phase 1 – energy modelling and analysis of the building the future sll building in fribourg will be constituted of 4 four floors and a basement for a total surface of about 5 000 m2 of floor area. it will serve as a hybriduse building, with offices and laboratory-type research area covering an energy reference area of 3 859 m2. the complete structure will be utilized as a research platform for conducting studies in real-world settings. the building is fitted with a double-flow mechanical ventilation system with heat recovery, and heat is provided via the air handling unit (ahu) connected to the neighbourhood’s district heating/cooling system using two heat exchangers. a number of radiant panels on the ceiling and a series of clay composite panels in the interior areas add to the building’s thermal inertia. the latter counterbalances the building’s low inertia as, except for the stairwells and basement, which are built of reinforced concrete the rest of the building is made in a light timber frame. domestic hot water (dhw) is created utilizing decentralized electric heaters installed at the consumption end points due to the low consumption of dhw and the need for maximum flexibility. this adaptability will help the structure and its installations evolve, as well as enabling the researchers’ tests to be carried out. an extensive research was done during the project’s design phase in order to employ the best-oriented and productive pv surfaces on the building exterior. the sll building will have a multi-oriented installation of 135 kwp (976 kwh/kwp). figure 1 shows the solar energy model built with rhino [11] and diva for grasshopper [12]. 2.2. phase 2 – design of a hydrogen-based storage system the h2 storage system is composed of an electrolyser, storage tanks and a fuel cell. on the electrical side, the h2 storage system works just like a battery pack, it allows to store electrical energy from the pv panels in the form of hydrogen and reuse when needed. on the thermal side, heat can be recovered from the fuel cell and stored in a hot water tank before being redistributed to the heating system. in parallel to the fuel cell, a small heat pump can cover the additional thermal needs when no hydrogen is available. on the far left of figure 2, an air-water heat exchanger is needed in the case where the fuel cell is working but no thermal demand is present. as it is presented in figure 2, in our case study, the main heating of the building is provided by the district heating, heat recovery from the fuel cell will be used for additional demand that consists of the heating of the basement. 282 vol. 38/2022 building-integrated photovoltaics (bipv) combined . . . figure 1. cumulative annual irradiation of bipv installation. figure 2. schematic design of the fuel cell in the heating system. the choice here was made to recover heat only on the fuel cell (not on the electrolyser) as it’s mainly used in the wintertime when the thermal demand is much higher. the simulation program integrates the following algorithm of energy management: each hour of the year is treated individually, when the pv production is higher than the demand, energy can be stored in the battery, used to create hydrogen or injected into the grid. when the demand is more important than the production electricity can be withdrawn from the battery or from the h2 storage system otherwise it can be imported from the grid. the program lets the user choose all the specifications regarding the h2 storage system, the battery pack and heat recovery on the fuel cell. the program then runs a simulation over a one-year period and presents the following output values that can be used to analyse the results. 2.3. phase 3 – sizing and optimisation process in this phase, the sizing of the hydrogen storage system is made for the sll. to run the simulations, the pv production, electrical consumption and thermal demand for the heating of the basement are added. two parameters can be tweaked in order to modify the simulations. firstly, the components that will be implemented in the building. for example, the volume of the h2 tanks, the capacity of the battery pack, the volume of the hot water tank or the electrical/thermal efficiency of the fuel cell. secondly, the way in which the system is controlled (control mode). that is, in our case, the way in which the fuel cell and electrolyser are activated. to size the system, the base mode has been used, meaning that the electrolyser will be turned on when the production is higher than the demand and the battery pack is fully charged. the fuel cell is turned on when the consumption is higher than the demand and the battery pack is empty. in a second step, using the same system other modes will be tested. four modes were developed specifically to achieve different goals: • maximum self-sufficiency ratio. • lowest injection/withdrawal peaks into and from the grid. • lowest ghg emissions. • maximum overall efficiency of the h2 storage system. in order to run multiple simulations, a vba code has been added to the program. it offers the opportunity to automatically test multiple scenarios by varying the input values and writing the output values in a .csv file. the data will then be analysed with the help of design explorer [10]. it is then possible to choose the best inputs to obtain the targeted results. in our case study, four inputs were chosen to be sized. the four components and the values chosen are as 283 s. aguacil, y. morier, p. couty, j.-p. bacher acta polytechnica ctu proceedings figure 3. sized system from design explorer. follow: • capacity of the h2 storage system (2 000, 4 000, 6 000, 8 000, 10 000, 12 000 [kwh pci]). • capacity of the battery pack (10, 20, 30, 40, 50 [kwh]). • volume of the hot water tank (1 000, 2 000, 3 000, 4 000, 5 000, 6 000, 8 000 [litres]). • efficiencies of the fuel cell: (1.) electrical 50 %, thermal 35 %; (2.) electrical 37 % thermal 43 %. by combining these inputs together, 420 scenarios are computed. to compare all the scenarios, we will look at three output values: the ss ratio, the amount of energy consumed by the additional heat pump and the overall thermal losses of the system. figure 3 shows the 420 solutions presented in design explorer. on the horizontal axis are the 5 input values and 3 outputs. on the vertical axis are the chosen values for the input and the obtained values for the outputs. the blue areas show all the solutions that allow to maximize the ss ratio and lower the heat losses and amount of energy used by the additional heat pump. finally, the green line shows the chosen solution that takes into account the limitations on the outputs and selects the smallest h2 storage, battery pack and hot water tank possible. 3. results in this section the main results are presented. 3.1. optimized sizing of the hydrogen storage system the sizing of the hydrogen storage system was made according to phase 3 of the methodology. using this process, the following values were found: • hydrogen storage system: 10 000 kwh pci (gross calorific power). • li-ion battery: 30 kwh. • hot water tank: 3 000 litres. • fuel cell: 50 % electrical efficiency and 35 % thermal efficiency. a 10 000 kwh pci storage capacity corresponds to 14.3 cubic meters of hydrogen stored at 300 bars or about 300 kg. the li-ion battery was sized to be as small as possible, prioritizing h2 storage. it’s worth mentioning that if the building had to reach the same ss ratio without h2 storage, the li-ion battery would need to be 17 times larger. the hot water tank also serves as an energy storage device. the bigger the tank, the more energy from the fuel cell it can store but as the tank gets bigger, the heat losses increase too. in our case, 3 000 litres appeared to be the ideal size. in this paper, two different fuel cells were considered, one with higher electrical efficiency and the other one with higher thermal efficiency. as it can be seen in figure 3 the only suitable fuel cell to achieve the goals fixed in the previous section is the one with the higher electrical efficiency (50 % electrical efficiency). for all further analysis this sized system will be used. 3.2. output values from the program in this section, the sized system will be used to present different results. all the following solutions use the same system but the control mode used will vary. the control mode is the way in which hydrogen production/consumption is managed. in the previous section, the base mode was used, meaning that the electrolyser and fuel cell are activated whenever it’s possible. this mode allows for the highest ss ratio but also means that the h2 storage is completely empty during a few months in the winter and entirely full during around two months in the summer (figure 4). the maximum ss ratio mode is the one that has been used so far and in our previous researches. the goal for other control configurations is to focus not only on the ss ratio but to show the other advantages of a h2 storage system. the second control mode aims to reduce production and consumption peaks (injection and withdrawal into the grid). one of the main issues with pv panels is the peak production happening around midday when the sun’s intensity is at its highest. the energy that cannot be used or stored is to be injected into the grid. this could soon 284 vol. 38/2022 building-integrated photovoltaics (bipv) combined . . . figure 4. amount of h2 stored over a one-year period time using the max ss ratio mode. figure 5. amount of h2 stored over a one-year period time using the peaks limitation mode. become a problem as more and more pv systems are installed and the grid would have to withstand a huge amount of energy over a short period of time. in this case, the h2 storage system limits the injection peaks by turning on the electrolyser only when the production exceeds a certain value. by comparing figures 4 and 5 we can see how the curve gets much smoother and the h2 storage tanks are almost never empty or full. the lowest ghg emission mode uses an hourly step-time data file containing the co2eq emissions factors for the swiss electricity mix over one year [13]. using this data, the h2 program can choose when to buy/sell electricity to the grid according to its carbon content. by withdrawing electricity from the grid only when the carbon content is low, the ghg emissions of the building related to electricity import can be kept at their lowest. the last mode aims to maximize the overall efficiency of the h2 system. the efficiency of the h2 storage system can be computed using equation (2). ηglobal = ηelectrolyser · qhydrogen qt hermal + qelectrical (2) equation (2) clearly shows that to maximize efficiency, both thermal and electrical energy needs to be used. the aim of this mode is to make sure the fuel cell is turned on only when there is both a thermal and electrical demand. to do so, the program will look at the temperature of the hot water tank and the electricity consumption of the building. if the temperature is not maxed out and there is an electricity demand, then the fuel cell will be activated. the results obtained with these different modes can be seen in table 1. 4. discussion in a previous study [5], a similar system was modelled without heat recovery, the overall efficiency of the hydrogen cycle back then was around 30 %. this value can be computed using equation (1) with qt hermal equalling zero. by recovering heat on the fuel cell, the efficiency can be increased by 65 % percent to reach 49.5 %. this clearly shows that to take full advantage of a hydrogen storage system, thermal recovery must be integrated. looking at the ss ratio, the sized h2 storage system allows to gain around 12 % reaching 80.7 %. however, it has been shown in previous computations that to further increase this ratio the amount of hydrogen needed is impossible to store. it would then become necessary to look at other factors such as the cost and carbon footprint required to build a space dedicated to the h2 storage tanks. regarding the ghg emissions, the chosen sized h2 storage system can help reach the same emissions but can’t improve this value. this is due to two factors. first, the ghg emissions related to the implementation of an electrolyser, storage tanks and fuel cell are still relatively high as these components are not very common. secondly, to the fact that in our computation, when electricity is exported from the building into the grid, the ghg emissions of the pv panels and the grid carbon content are compared. as the carbon content of the electricity from the pv system is always lower than the grid’s, for every kwh exported a negative value is added to the ghg emission balance. when electricity is stored and reused, the model does not take advantage of this benefit. if we were to compute ghg emission balance without this deduction, the h2 storage system would always have a better carbon footprint. this h2 storage system also allows to reduce both injection and consumption peaks. with the dedicated mode, the injection peaks drop from 143 to 99 kw and consumptions from 51.2 285 s. aguacil, y. morier, p. couty, j.-p. bacher acta polytechnica ctu proceedings mode ss ratio ghg emissions overall max withdrawal max injection[%] [kg co2eq] efficiency [%] [kw] [kw] max ss ratio 80.7 9671 34 47.5 140 peaks limitation 72.5 6764 48.5 39 99 max efficiency 77 8071 49.5 45.3 140 ghg limitation 73.6 5658 41 47.5 142.8 no h2 storage 69 5586 0 51.2 142.8 table 1. results from the different program modes. to 39 kw preventing overloads on the network. this could also be very beneficial to a neighbourhood like bluefactory [14] in fribourg, where the sll will be built, as it would be possible to store or distribute energy from/to all the buildings in the area. it should be remembered that all these results are obtained using different control modes and that just one wouldn’t be able to reach all of these values at once. the financial aspect of this h2 storage system still needs be evaluate in further details. it’s is still complicated to estimate prices regarding the components of an h2 storage system. moreover, with today’s electricity prices in switzerland [15] the building would basically have no electrical cost as the amount of energy sold to the grid is much higher than what’s consumed. the constant evolution of electricity prices and the feed-in tariffs should be also taken into account. further research will be necessary to evaluate every aspect of an h2 storage system. financial facets still need be analysed in details to evaluate in which context this installation would be profitable. the h2 program could be optimized by selecting the best mode for every case or by writing other modes allowing to furthermore improve system operation. the h2 storage system could be used not only for the building but in harmony with the grid allowing, when needed, to store energy peaks or to provide great amount of energy. new ideas could also be implemented in this project with the potential of hydrogen vehicle. a quick calculation shows that the sll would have the potential to produce enough h2 for cars to travel about 150 000 km each year. 5. conclusions this paper demonstrates the many interesting aspects of integrating a hydrogen-based electricity storage system at building-scale. this case study shows the potential of h2 storage and the various benefits that it could have. the use of thermal energy released by the fuel cell is necessary to greatly improve the overall efficiency of the h2 cycle. to pursue this project, the ms excel written code will be improved and optimized, further analysis will be conducted to evaluate the financial aspects in more details and new opportunity for the use of h2 will be implemented. the challenges regarding energy storage will become more and more common and this paper clearly demonstrate that h2 will have a role to play in the future of renewable energy. acknowledgements the authors thank the university of applied sciences and arts western switzerland (hes-so) and the ecole polytechnique fédérale de lausanne (epfl) for their support. references [1] t. hostettler, a. hekler. statistik sonnenenergie. referenzjahr 2020, bundesamt für energie bfe, 2021. [2] a. parlikar, c. n. truong, a. jossen, h. hesse. the carbon footprint of island grids with lithium-ion battery systems: an analysis based on levelized emissions of energy supply. renewable and sustainable energy reviews 149:111353, 2021. https://doi.org/10.1016/j.rser.2021.111353 [3] p. couty, d. torregrossa, j.-p. bacher. battery storage systems for building applications. energy in buildings and communities programme 2020. [2021-05-24]. https://arodes.hes-so.ch/record/8139 [4] p. couty, m. j. lalou, p. cuony, et al. positive energy building with pv facade production and electrical storage designed by the swiss team for the u.s. department of energy solar decathlon 2017. energy procedia 122:919–924, 2017. https://doi.org/10.1016/j.egypro.2017.07.410 [5] y. morier, s. aguacil moreno, p. couty. hydrogenbased electricity storage optimization on buildings by coupling thermal and photovoltaic electricity production towards carbon neutrality. journal of physics: conference series 2042(1):012092, 2021. https://doi.org/10.1088/1742-6596/2042/1/012092 [6] vela solaris ag. polysun software, 2018. [2021-05-24]. https://www.velasolaris.com/ [7] smart living lab. smart living lab, 2021. [2021-05-24]. https://www.smartlivinglab.ch [8] designbuilder. designbuilder software v.7, 2021. [2021-01-29]. http://www.designbuilder.co.uk/ [9] u.d. of e. doe, us department of energy, 2018. [2018-04-15]. https://www.energy.gov/ [10] core studio, t. thornton. design explorer v.2 for multi-dimensional parametric studies, 2017. [2018-0509]. http://tt-acm.github.io/designexplorer/ [11] robert mcneel & associates. rhinoceros software, 2019. [2019-02-01]. https://www.rhino3d.com/ 286 https://doi.org/10.1016/j.rser.2021.111353 https://arodes.hes-so.ch/record/8139 https://doi.org/10.1016/j.egypro.2017.07.410 https://doi.org/10.1088/1742-6596/2042/1/012092 https://www.velasolaris.com/ https://www.smartlivinglab.ch http://www.designbuilder.co.uk/ https://www.energy.gov/ http://tt-acm.github.io/designexplorer/ https://www.rhino3d.com/ vol. 38/2022 building-integrated photovoltaics (bipv) combined . . . [12] solemma lcc. diva for rhino. environmental analysis for buildings, 2018. [2018-04-15]. http://diva4rhino.com/ [13] d. vuarnoz, s. aguacil moreno. dataset concerning the hourly conversion factors for the cumulative energy demand and its non-renewable part, and hourly ghg emission factors of the swiss mix during a one-year period (2016 and 2017). data in brief 30:105509, 2020. https://doi.org/10.1016/j.dib.2020.105509 [14] bluefactory. bluefactory – fribourg – freiburg sa, 2014. [2021-05-24]. https://www.bluefactory.ch/ [15] vese. feed-in tariffs for pv solar energy, 2018. [2019-02-01]. http://www.vese.ch/ 287 http://diva4rhino.com/ https://doi.org/10.1016/j.dib.2020.105509 https://www.bluefactory.ch/ http://www.vese.ch/ acta polytechnica ctu proceedings 38:281–287, 2022 1 introduction 2 methodology 2.1 phase 1 – energy modelling and analysis of the building 2.2 phase 2 – design of a hydrogen-based storage system 2.3 phase 3 – sizing and optimisation process 3 results 3.1 optimized sizing of the hydrogen storage system 3.2 output values from the program 4 discussion 5 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0057 acta polytechnica ctu proceedings 38:57–64, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague towards an automatized and objective assessment of data from visual inspections of building envelopes jan mandinec∗, pär johansson chalmers university of technology, department of architecture and civil engineering, sven hultins gata 6, 412 58 gothenburg, sweden ∗ corresponding author: jan.mandinec@chalmers.se abstract. the renovation planning process is filled with uncertainties and subjective decisions. these make the decisions upon what and when to renovate a complex and ambiguous problem. selection of renovation measures related to building envelope are often far from optimal as decisions are usually made based on visual inspections. these are manned and thus prone to subjective assessment and the knowhow of individual inspectors. furthermore, objective criteria which could indicate non-structural failures are often missing. the objective based planning process allowing the estimation of the current damage status of the building envelope by only using non-destructive measurements is still in its infancy. the first step requires establishing reliable and objective based data collection. these could be efficiently collected by unmanned aerial vehicles (uav) with subsequent image recognition algorithms allowing the identification of imperfections and store the position and extent of such deviations into the building’s digital assessment database. such tools do not exist. the aim of this study is to investigate the current objectivization possibilities in the domain of building inspections. the first part provides a literature review describing how an autonomous uav survey of a building envelope may be planned and what computer vision techniques may be used for automatic damage recognition and classification. subsequently, an objective detection model based on the yolo-tiny (you only look once) computer vision framework is employed in a case study investigating a building envelope of historical tjolöholm castle in sweden. this study contributes to developing a methodology for an objective based visual inspection process. keywords: uav, computer vision, object detection, yolo, case study, building, renovation, brick, masonry. 1. introduction in terms of environmental aspects, renovation of a building or its adaptive reuse has proven to outperform constructing a new building [1]. the environmental effectiveness of renovations may be further improved by transforming the maintenance paradigm from reactive (emergency renovation after a component breaks down) to proactive (planned renovation before a component breaks down). however, this paradigm shift requires a full digitalization of the maintenance planning allowing to perform data driven prediction of failures (predictive maintenance). inspiration could be taken from the world of wind turbines, or jet fighters where data can already predict whether any component will fail within a given period allowing not only to reduce down-times but also optimize the usage of resources and thus the environment. ideal maintenance system in buildings should be capable of doing the same type of assessment to recommend measures which can prevent a failure. however, reaching this level of predictive maintenance requires systematic changes in data collection processes. the performance of machine learning algorithms allowing a computer to find patterns in the data and thus automatically learn and decide over complex systems, are highly dependent on the type and the total amount of the data. building inspection provide invaluable information for future failure prediction as it shows how much damaged building is at the time of inspection. however, common manned inspections are insufficient in terms of large data collection. moreover, data are often not stored in a format suitable for applying machine learning algorithms. note that insufficient in this context means that building inspectors, however experienced, cannot fully quantify every aspect of how a building envelope is damaged (e.g., count the number of all cracks, measure, and sum up all damaged areas). contrastingly, unmanned aerial vehicles (uavs or drones) have such possibilities as drones can gather large sets of data by scanning whole building envelope with various sensors (i.e., rgb or infrared cameras, lidar, radar). given the premise from above, one need to objectively obtain the current status of the building envelope. the most readily available sensor technique is represented by rgb (red, green and blue) cameras. the overarching research questions this study answers are: 57 https://doi.org/10.14311/app.2022.38.0057 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en jan mandinec, pär johansson acta polytechnica ctu proceedings (1.) which data can objectively be used to assess what the current and future performance level of a building envelope are? (2.) how can a damage on a building envelope be quantified based on images from uavs? (3.) how can the whole process become as automatized as possible? this paper is divided into a literature review and the case study. the literature review is further divided into two distinct parts. based on the conclusions from the literature review, the case study is used to optimize the methodology into the practice. this is done by applying the in-house build sense model, developed based on the yolo-tiny (you only look once, tiny version) object detection framework [2], on a wall of tjolöholms castle in the west of sweden. finally, the study concludes with the discussion about how damaged areas may be quantified giving the outlook for future data-driven uav based building inspections. 2. literature review the first part of the literature review investigates how a flight pattern of uavs may be automatically generated for building envelope inspection. the second part presents automatic damage recognition from images to investigate how damages on a building envelope may be automatically recognized by the use of computer vision algorithms. 2.1. planning of uav based inspections of building envelope usage of unmanned aerial vehicles (uavs or drones) in a building inspection process yields many advantages compare to the classical manned inspections. the two of the most distinct advantages are the time efficiency and the reachability over the whole building envelope. the subsequent analysis of the images captured by drones may provide a comprehensive look on the status of the building envelope. the building inspection process based on a uav survey may be either conducted manually i.e., the drone is controlled remotely by an operator on the ground, or it can be conducted semi-automatically. in both cases local laws and restrictions regarding operations of the uavs need to be followed. duque [3] identified 4 necessary steps which need to be taken prior to the inspection of any construction: (1.) uav fly operation training, (2.) documentation review of inspected structure, (3.) observation of surroundings and (4.) uav pre-flight check (e.g., battery levels). in the latter semi-automatic case, drone is often preprogramed to follow specific gps coordination waypoints and flight path making the flight itself fully automatic. such an approach was adopted e.g. in a study by rakha [4] who used images captured by infrared camera to develop heat-leakage segmentation framework based on computer vision algorithms. however, the loss of the gps signal may lead to navigational hazards. to cope with this problem, a slam (simultaneous localization and mapping) technique, was developed, allowing drones to detect objects in the surroundings by employing lidar or similar sensor and thus to avoid collisions. current development in the semi-autonomous flight systems adopts 3d models (or bim models) of an inspected building as a source of prior information for uav inspection planning. freimuth [5] utilizes bim models to automatically generate safe flight paths for uavs. similar framework developed for automatized inspections of bridges, allowing to generate safe flight paths can be found in the work by morgenthal [6]. this framework was practically utilized for inspecting building envelopes by benz [7] who performed a case study to automatically generate flight paths for obtaining infrared images and subsequent thermal conductance value determination of the building envelope. note that flight path in this case needed to be manually adjusted to avoid collision with nearby obstacles (e.g. trees). regardless the type of flight control, uavs need to employ specific routing strategies to comply with the needs of different sensors and the intentions of the quest. for this purpose, a different flight pattern is needed for drone scanning to produce high resolution images suitable for subsequent defects analysis compared to a drone taking infrared images suitable for thermography analysis. 2.2. automatize damage recognition and classification from images recent development of computational power made the machine learning and computer vision algorithms accessible to the point where increasing numbers of research papers dealing with their application on building inspections are being published every year. the foundation for the majority of algorithm for image interpretations is a convolutional neural network (cnn) as it mainly specializes on extracting basic features from images like straight lines, round shapes etc. extracted features are then subsequently fed to a classification algorithm providing the predictions. the aforementioned method was successfully used for crack detection where support vector machine (svm) algorithm achieved 85.94 % of the validation accuracy [8]. note that substantial number of labelled images i.e., images where the labels are specified manually, are needed for reaching high levels of accuracy. the aforementioned study used 6 002 images of “cracks” and “non-cracks”. more examples of similar detection strategy using different classification algorithms can be found in a literature review conducted by sony [9]. an interesting option allowing to overcome the need of the high number of images is represented by the concept of transferred learning which is the 58 vol. 38/2022 towards an automatized and objective assessment . . . domain of computer vision frameworks e.g., r-cnn (region-based cnn), faster r-cnn, alexnet, oxfornet, yolo (you only look once), resnet50, vgg16. such frameworks are usually based on layer(s) of cnns subsequently connected to feed-forward neural networks or similar for classification purposes and in case of object detection problems also to a regression algorithm, which also allows to accurately draw a bounding box around the objects area. note that frameworks are obtainable as pre-trained i.e., the user doesn’t need to train the whole network from scratch but rather tune it for a specific problem. in some specific cases, this may lead to the reduction of the total number of images needed for achieving accurate predictions. özgenel [10] compared the performance of 7 different computer vision frameworks and concluded that transfer learning applied to crack detection problems is feasible. in terms of transfer learning, the study also advised to use limited number of training images especially when the variance among the images is low as the high number of images increases computational time as well as the risk that the model learn some specific crack patterns too well and fails to comprehend general appearance of cracks (overfitting). the majority of research papers investigating the usage of presented computer vision techniques comes mainly from three research areas: bridge health monitoring, pavement monitoring and the large-scale monitoring of buildings after earthquakes. the first two areas use the techniques described above for damage detection. the latter mainly uses 3d cloud of points (set of points in virtual 3d space usually forming a base for 3d models) to quickly find collapsed buildings and subsequently applies computer vision techniques to detect damages of smaller scale [11]. practical example for concrete bridges is provided by yang [12] who performed a field test with a drone and by the means of vgg-16 framework successfully detecting over 70 % of cracks and spalling areas on a ground level of a bridge. the means of transferred learning was used to detect cracks on pavements in the work of gopalakrishna [13] who used 1 056 images in total for training purposes. more computer vision based examples of bridge health and pavement monitoring alongside with examples of different applications (building health monitoring or inspection of underground structures) may be found in the work of sony [9]. practical applicability of pre-trained frameworks in detecting defects on building envelopes may be well illustrated on the work done by wang [14] who feed 500 annotated images to faster r-cnn framework to identify and draw bounding boxes around bricks subjected to spalling and efflorescence with resulting accuracy 95 %. another good example is provided by the research of cha [15] where modified faster r-cnn is used to differentiate concrete cracks, steel corrosion, bolt corrosion and steel delamination. in total 2 366 annotated training images were fed to the framework achieving mean average precision 89.7 % while testing on previously unseen images. image segmentation is another computer vision technique allowing to precisely highlight the area of interest i.e., to color out pixels which are associated with a crack or other type of damage. hoskere [16] used a cnn on a level of individual pixels rather than on whole images eventually adding a label to each pixel and thus highlighting six different types of damage. already mentioned work by gallareta [11] combined multiresolution segmentation allowing to differentiate building objects (i.e., façades, windows etc.) and spectral difference segmentation allowing to highlight damaged areas (i.e., cracks, holes etc.). outputting highlighted areas from image segmentation may be used for measuring distances within an image eventually allowing to measure the scope of a damage. such approach was illustrated in the work of lins [17] who used images to measure width and length of cracks. 3. case study in the following case study, an in-house computer vision algorithm (build-sense model) based on the yolo-tiny framework [2] capable of detecting damaged areas is applied to images of historical tjolöholm castle in the west of sweden. the main goal is to detect missing joints in a stone masonry wall. 3.1. object detection model the build-sense model was originally built for detecting damage on brick walls specializing on drawing bounding boxes around cracks and bricks subjected to spalling and efflorescence. the model is utilizing the concepts of transfer learning and was trained based on 225 images of damaged brick walls (approximately 75 images for each damage type) manually annotated with more than 2 000 bounding boxes providing the model with the interpretation of how a damage looks like. figure 1 shows bounding boxes around efflorescence as they were predicted by build-sense model. the current predictive power of the build-sense model may be characterized as good enough for initial stage of inspections as it in theory may notify the inspector about possible building envelope failure. however, given its rough estimation of the damaged area it is not yet suitable for accurate damage quantification i.e., measuring the area of damage, width of damage etc. 3.2. tjolöholms castle and data collection the construction of tjolöholms castle was finished in the beginning of 20th century. thenceforth, it has been subjected to many refurbishments given its continuous moisture related problems which occurs in the form of mold growth and ongoing disintegration of joints in the stone masonry walls forming the building envelope of the castle. problems are more obvious on 59 jan mandinec, pär johansson acta polytechnica ctu proceedings figure 1. example of output from the build-sense model – model draw bounding boxes around damaged area alongside with its level of confidence [18]. the façade facing the south-east with the open view towards the coast of north sea. as foundation for the status assessment a basic ground level visual inspection of the building envelope was performed and complemented by inspecting building insides, including the roof. following the many years of subsequent repair work in the castle, there is a substantial amount of data. this is stored in the castle management’s own databases containing technical drawings, data from hygrothermal wall sensors, 3d model of the building (based on cloud of points), panoramic images of each room from interior and panoramic images taken on 78 different locations around the main building. note that performing an uav survey for acquisition of close-up images of building envelope was not possible at the time of the visit. therefore, obtained outdoor panoramic images are used in subsequent computer vision analysis. 3.3. scope and the methodology of the case study given the fact that joints are continuously disintegrated, mainly on the east side of the south-west façade, the study is limited to the assessment of the build-sense model only to one panoramic image which closely depicts the area of interest. the positions of the panoramic image as well as the overall view over the wall of interest is on the figure 2. given the limitations of the position i.e., testing images will not be in direct opposition to the wall surface but rather tilted, and given the complexity of light exposure conditions (uneven surface is under direct sunlight casting a lot of shadows), one area of interest on a building envelope will be analyzed from different angles. the confidence threshold for the model was set to be 50 % i.e., the model draws bounding boxes around objects only where the probability of accurate detection is higher than 50 %. 3.4. retraining the build-sense model direct application of the current build-sense model is not feasible in this problem as it was not able to detect any missing joints in the castle’s stone masonry wall. this is expected as it was originally tuned for detecting cracks, efflorescence and spalling on brick walls i.e., each construction has different structure and different damage patterns which are different from missing joints in stone masonry walls. therefore, the model needed to be retrained specifically for the purpose of detecting missing joints in stone masonry walls. images for re-training purposes were collected in the streets of gothenburg city (in the west of sweden) by the in-built high-resolution camera in iphone se 2019. initial 115 images of stone masonry walls were annotated highlighting damaged joints. the total number of images was multiplied in the image augmentation process by performing 90°, 180° and 270° rotation. the images were then split to train/test samples by the ratio of 80/20. training itself was performed in the google collaboration environment allowing to borrow graphics processing units (gpu) for the purpose of calculation. the whole training session took 6.3 hours to complete on a cloud based dedicated gpu nvidia tesla k80 controlled by a local computer with 2,.3 ghz quad-core intel core i5 processor. note that the time of the calculation could been decreased by limiting the pixel size of images used for the training purposes. the model iterated over the images for 6 000 times internally using mean average precision (map) as a metric for network’s weights evaluation. final output of the training is a set of weights which are the best in predicting missing joints given the training images and the number of iterations. the best weights are then used for subsequent analysis of images from tjolöholms castle. 60 vol. 38/2022 towards an automatized and objective assessment . . . (a). (b). figure 2. (a) schematic plan on the building’s ground floor to the left shows the location where the panoramic photograph was taken (red dot). gray background indicates multi-storey parts while cyan colour indicates single storey part of the building. (b) photograph was taken on the stairs leading to the roof of single storey part [19]. figure 3. default view from the panoramic image directly facing the area of interest [19]. 3.5. results and discussion all images altogether presented in this subchapter were analyzed using retrained build-sense model. figure 3 shows the default (un-zoomed) view from the panoramic image directly facing the building envelope. note that there are missing joints on the left side of the window. the model there fails to detect any objects of interest. however, this was not a case for zoomed images. figure 4a shows missing joints in the vertical direction but the algorithm failed to comprehend damages in the horizontal direction. observe that the confidence of the model in predicting missing veneer is relatively high i.e., always above 70 %. complete opposite picture gives the figure 4b where the missing joints were detected mainly in horizontal direction while the model’s confidence is in general lower. the model failed to comprehend a majority of missing joints in figure 4c. however, vertical joints are in this point of view badly visible and difficult to correctly interpret even for human eyes. furthermore, the model again fails to comprehend joints in horizontal direction which makes the confidence of the model over individual bounding boxes rather low. observe that the model prefers detecting missing joints of relatively thin size but fails to detect thicker joints. this is probably caused by the limited number of training images and by the low variance of images i.e., training set does not include missing joints of this size. the presumption that the low confidence in the previous case is caused by the low number of relevant training images is further strengthened by figure 5 where most missing joints were correctly detected with high level of confidence. the model only fails to grasp missing veneers in the opposite wall in the distance and few parts of the joints on the sides. 61 jan mandinec, pär johansson acta polytechnica ctu proceedings (a). detailed view 1. (b). detailed view 2. (c). detailed view 3. figure 4. detailed views. credit: anders jansson, all rights reserved. figure 5. image facing stairs [19]. 62 vol. 38/2022 towards an automatized and objective assessment . . . 4. conclusion this study investigated how the inspection process of a building envelope may be automated by the means of uavs and subsequent computer vision based analysis of images. the first part of the paper is a literature review summarizing how drones may be automatically navigated to capture data from a building envelope. the second part of the literature review shows how computer vision algorithms are used in image analysis. the latter part of the review was exemplified using computer vision based model (build-sense model) for detection of missing joints on a wall of tjolöholms castle, a historical building in sweden. the model failed to comprehend majority missing joints in vertical direction and joints of larger thicknesses. it was concluded that the model may be improved by adding a variance to the training set (more different examples) and by expanding the total number of the training set. nevertheless, current predictive power of the model may be typified as good enough for notifying inspectors, who should be responsible for final damage evaluation, about potential building envelope failure. despite the performance limits of the presented model, the conjunction of fast and overreaching uavs with the computer vision promises to bring revolution to the building inspection process. when properly set, uavs may scan whole building envelope in terms of minutes. subsequent analysis may be used for accurate damage quantification of building envelope i.e., detecting the number of cracks per square meter, measuring the scope of a damage etc. by separating the envelope into smaller parts, one could differentiate the performance of the envelope over those parts and subsequently use this information for accurate data-based failure predictions, opening the gate of predictive maintenance. acknowledgements this work was supported by the swedish research council for environment, agricultural sciences and spatial planning (formas) [grant no. 2019-01402]. the authors also thank chalmers information and communication technology (ict) area of advance for additional financial support allowing the development of the build-sense model. references [1] v. hasik, e. escott, r. bates, et al. comparative whole-building life cycle assessment of renovation and new construction. building and environment 161:106218, 2019. https://doi.org/10.1016/j.buildenv.2019.106218 [2] z. jiang, l. zhao, s. li, y. jia. real-time object detection method based on improved yolov4-tiny, 2020. 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[13] k. gopalakrishnan, s. k. khaitan, a. choudhary, a. agrawal. deep convolutional neural networks with transfer learning for computer vision-based data-driven pavement distress detection. construction and building materials 157:322–330, 2017. https: //doi.org/10.1016/j.conbuildmat.2017.09.110 [14] n. wang, x. zhao, p. zhao, et al. automatic damage detection of historic masonry buildings based on mobile deep learning. automation in construction 103:53–66, 2019. https://doi.org/10.1016/j.autcon.2019.03.003 [15] y.-j. cha, w. choi, g. suh, et al. autonomous structural visual inspection using region-based deep learning for detecting multiple damage types. computeraided civil and infrastructure engineering 33(9):731– 747, 2018. https://doi.org/10.1111/mice.12334 63 https://doi.org/10.1016/j.buildenv.2019.106218 https://doi.org/10.48550/arxiv.2011.04244 https://doi.org/10.1061/(asce)be.1943-5592.0001289 https://doi.org/10.1061/(asce)be.1943-5592.0001289 https://doi.org/10.1080/24751448.2018.1420963 https://doi.org/10.1016/j.autcon.2018.10.016 https://doi.org/10.1016/j.autcon.2018.10.006 https://doi.org/10.1016/j.enbuild.2021.111266 https://doi.org/10.21660/2018.51.35376 https://doi.org/10.1016/j.engstruct.2020.111347 https://doi.org/10.22260/isarc2018/0094 https://doi.org/10.5194/nhess-15-1087-2015 https://doi.org/10.1016/j.conbuildmat.2017.09.110 https://doi.org/10.1016/j.conbuildmat.2017.09.110 https://doi.org/10.1016/j.autcon.2019.03.003 https://doi.org/10.1111/mice.12334 jan mandinec, pär johansson acta polytechnica ctu proceedings [16] v. hoskere, y. narazaki, t. a. hoang, b. spencer. vision-based structural inspection using multiscale deep convolutional neural networks, 2018. [2022-02-14]. https://doi.org/10.48550/arxiv.1805.01055 [17] r. g. lins, s. n. givigi. automatic crack detection and measurement based on image analysis. ieee transactions on instrumentation and measurement 65(3):583–590, 2016. https://doi.org/10.1109/tim.2015.2509278 [18] karagrubis. brick efflorescence. https://rainguardpro.com/ [19] a. jansson. tjolöholms castle. 64 https://doi.org/10.48550/arxiv.1805.01055 https://doi.org/10.1109/tim.2015.2509278 https://rainguardpro.com/ acta polytechnica ctu proceedings 38:57–64, 2022 1 introduction 2 literature review 2.1 planning of uav based inspections of building envelope 2.2 automatize damage recognition and classification from images 3 case study 3.1 object detection model 3.2 tjolöholms castle and data collection 3.3 scope and the methodology of the case study 3.4 retraining the build-sense model 3.5 results and discussion 4 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0027 acta polytechnica ctu proceedings 40:27–32, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague modelling of laminated glass interlayer by fractional viscoelasticity barbora hálková∗, jaroslav schmidt czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: barbora.halkova@fsv.cvut.cz abstract. fractional calculus, i.e. the theory of derivatives and integrals of non-integer order, can be efficiently used for the theoretical modelling of viscoelastic materials. our research is focused on the polyvinyl butyral which is used as an interlayer for the laminated glass. polyvinyl butyral can be classified as a viscoelastic material and the introduction of the fractional viscoelasticity seems to be appropriate tool for its description. this paper briefly introduces the springpot element and its connection into more complex theoretical models. we mainly consider the generalized maxwell model in its standard and fractional form and show their application by fitting the data obtained by experimental analysis. keywords: laminated glass, polymer interlayer, theory of viscoelasticity, fractional viscoelasticity, springpot, generalized maxwell model, fractional calculus, storage modulus. 1. introduction laminated glass, see [1], is a composite material, which consists of a solid glass plates (float, toughened, tempered glass or their combination) and interlayers made of polymers (mostly the polyvinyl butyral, which will be also considered in the following text) or cast resins. the interlayer has primarily the safety function, it improves the post-breakage behaviour of the glass elements. it also ensures interaction between the main load-bearing glass plates, its effect is mostly important in the case of bending, when the interlayer directly affects the stiffness of the cross section and is mostly stressed by shear. the polyvinyl butyral (pvb) type of interlayer can be classified as a viscoelastic material [2]. its behaviour is somewhere between purely elastic and purely viscous. the viscoelastic behaviour is time and load-history dependent, phenomenon such as creep and relaxation are significant. for the accurate description of the viscoelastic material we need to use more complex theoretical models, while the ideal elasticity and viscosity stand for the limit cases, see [3] for broader description. 2. viscoelasticity the standard theory of viscoelasticity commonly uses the theoretical models composed of elastic and viscous elements (springs and dashpots). the fractional viscoelasticity, see [4], applies the principles of fractional calculus and introduces another rheological element, the springpot. however, the fractional calculus is not discussed in detail here because it is out of scope of this article and the interested reader is referred to [5]. e (a). spring. e, τc, α (b). springpot. η (c). dashpot. figure 1. rheological scheme of spring, springpot and dashpot. the linear spring (see figure 1a) is a rheological element, which represents ideally elastic behaviour of the material. the constitutive law, also known as hooke’s law, has the following form: σ(t) = eε(t), (1) where σ stands for the stress, ε for the strain, e for young’s modulus of elasticity. the stress is a linear function of the strain. the viscous damper (dashpot) on the other hand, the viscous damper (dashpot) (see figure 1c) represents the model of ideally viscous fluid. the stress is directly proportional to the strain rate (i.e. to the first derivative of the strain). the constitutive law, also known as newton’s equation of viscosity, has the following form: 27 https://doi.org/10.14311/app.2023.40.0027 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en barbora hálková, jaroslav schmidt acta polytechnica ctu proceedings σ(t) = ηε̇(t), (2) where η stands for the coefficient of viscosity, ε̇ the time derivative of the strain. the springpot (see figure 1b) then represents a transition between the two cases mentioned above. fractional viscoelasticity, see [4], assumes the constitutive law in the following form: σ(t) = ξdαε(t), (3) where ξ and α are parameters of the springpot, dα denotes the α-th time derivative. the point of fractional viscoelasticity is to assume α as non-integer. we have already mentioned that the ideal elasticity and ideal viscosity stands for the limit cases of viscoelastic behaviour, therefore we can define the limit cases for the springpot parameters and consider α only in the interval ⟨0, 1⟩. when we introduce the characteristic time τc in the following form: τc = η e , (4) equation (3) can be rewritten as: σ(t) = eτ αc d αε(t). (5) the replacement of ξ by eτ αc better illustrates the physical meaning of the constant ξ in the limit cases. for α = 0 we receive: σ(t) = e d0ε(t) dt0 = eε(t), (6) describing purely elastic response, remind equation (1). we see that in this limit the parameter ξ is equal to the modulus of elasticity e and the behaviour of the springpot corresponds to the behaviour of an elastic element. on the other hand, for α = 1 we receive: σ(t) = eτc dε(t) dt = ηε̇(t). (7) in this limit case the behaviour of the springpot corresponds to the behaviour of the viscous element (remind equation (2)), while the constant ξ become the constant of viscosity η. 2.1. relaxation modulus of springpot the relaxation is the phenomenon where the stress gradually decreases over time even if the strain remains constant. on the other hand, the creep can be explained as a phenomenon where the the strain gradually increases over time while the stress remains 0 20 40 60 80 100 time [s] 0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 re la xa tio n m od ul us [p a] spring =0,05 =0,2 =0,4 =0,6 =0,8 dashpot figure 2. relaxation modulus of a springpot according to different values of α. constant, for the broader description of these phenomena see [3]. in the following text we focus on the relaxation mode only, because it can be more easily interpreted in connection with the results of our experimental analysis. relaxation modulus of the material is defined as the stress response to the heaviside strain load: ε(t) = h(t), (8) unit-value strain load which is applied instantaneously in time t = 0 and remains constant for times t > 0. in the case of the springpot element we consider the fractional derivative dα as the riemann-liouville type of derivative, see [6]. the resulting relaxation modulus has the following form: r(t) = e γ(1 − α) ( t τc )−α h(t), (9) where γ denotes the gamma function. figure 2 shows the relaxation modulus of a springpot according to different values of α, while the dashed lines represent the limit cases given by elastic and viscous behaviour. 2.2. springpot under harmonic load as a part of our research the experimental analysis of the laminated glass is provided as well. the samples are loaded by a harmonic torque. therefore we are interested in the behaviour of theoretical models under the harmonic load as well. the prescribed strain load and the corresponding stress response have the following forms: ε∗(ω) = ε∗0e iωt, σ∗(ω) = σ∗0 e iωt, (10) where ε∗0 and σ∗0 are the amplitudes of the harmonic strain and stress in the complex plane, 28 vol. 40/2023 modelling of laminated glass interlayer by fractional viscoelasticity eη (a). maxwell model. ee, τc, α (b). fractional maxwell mode. figure 3. rheological scheme of the maxwell model and the fractional maxwell model. e is the euler constant, i is the imaginary unit. the complex modulus e∗ can now be defined as: e∗(ω) = σ∗0 ε∗0 . (11) if we separate the real part of the complex modulus, we receive the storage modulus e′ which corresponds to the elastic part of the response. on the other hand, the imaginary part, the loss modulus e′′ represents the viscous part of the response. in the following text we focus on the storage modulus due to better accuracy with which the storage modulus is obtained from experiments. the complex modulus of the springpot element is written as: e∗(ω) = e(iωτc)α. (12) the storage modulus is then obtained by separating the real part of the complex modulus as: e′(ω) = re(e∗) = e(ωτc)α cos ( α π 2 ) . (13) 2.3. theoretical models to improve the description of the viscoelastic behaviour, the rheological elements can be arranged together in series or in parallel to create more accurate models. the overview of some theoretical models can be found for example in [7], while here we limit attention to those, which seem to be appropriate for the description of the pvb interlayer. maxwell cell is a serial connection of an elastic spring and a viscous dashpot. by the replacement of the viscous element with the springpot we obtain the fractional maxwell cell. both rheological schemes with the variable parameters of each element are shown in figure 3. the relaxation moduli of the maxwell cell and the fractional maxwell cell are respectively: r(t) = ee−t/τc h(t), (14) r(t) = eeα,1 ( − ( t τc )α) h(t), (15) where e1 η1 e2 η2 en ηn e0 figure 4. rheological scheme of the generalized maxwell model. e1 e2 en e0 e1, α1 τc,1 e2, α2 τc,2 en , αn τc,n figure 5. rheological scheme of the fractional maxwell chain. eα,1 is the mittag-leffler function. for the limit case α = 1 it holds that e1,1(x) = ex, see [8], and therefore equation (14) and equation (15) become identical in this limit case. the storage moduli of the maxwell cell and the fractional maxwell cell are then: e′(ω) = e ω2τ 2c ω2τ 2c + 1 , (16) e′(ω) = e (τcω)2α + (τcω)α cos (α π2 ) (τcω)2α + 2(τcω)α cos (α π2 ) + 1 . (17) generalized maxwell model (also known as the maxwell chain) is the parallel connection of a spring and n maxwell cells, see figure 4. the fractional maxwell chain is then obtained as a parallel connection of a spring and n fractional maxwell cells, see the rheological scheme in figure 5. 29 barbora hálková, jaroslav schmidt acta polytechnica ctu proceedings 10 5 10 2 101 104 time [s] 0 20 40 60 80 100 re la xa tio n m od ul us [p a] spring + 1 maxwell cell spring + 3 maxwell cells spring + 5 maxwell cells figure 6. relaxation modulus of the maxwell chain compare to the number of cells. 10 5 10 3 10 1 101 103 angular velocity [rad/s] 106 107 st or ag e m od ul us [p a] fitted maxwell chain experimental data figure 7. fitting experimental data of g′ to 1-cell maxwell chain. the parallel connection of the cells in the generalized model enables us to obtain the relaxation modulus of the maxwell (resp. the fractional maxwell) chain superpositionally. the same applies to the storage modulus. the relaxation modulus and the storage modulus of the generalized models have the following form: r(t) = e0 + n∑ i=1 ri, (18) e′(t) = e0 + n∑ i=1 e′i , (19) where e0 is the young modulus of elasticity of the spring, ri are the relaxation moduli, e′i the storage moduli of each of the maxwell (resp. the fractional maxwell) cells. it is obvious that the behaviour of these generalized models depends on the number of the maxwell (resp. the fractional maxwell) cells connected together. this is illustrated in figure 6 which shows the difference in the relaxation modulus of a standard maxwell chain with 1, 3 and 5 connected maxwell cells. 10 5 10 3 10 1 101 103 angular velocity [rad/s] 106 107 st or ag e m od ul us [p a] fitted maxwell chain experimental data figure 8. fitting experimental data of g′ to 3-cell maxwell chain. 10 5 10 3 10 1 101 103 angular velocity [rad/s] 106 107 st or ag e m od ul us [p a] fitted maxwell chain experimental data figure 9. fitting experimental data of g′ to 7-cell maxwell chain. 2.4. validation of theoretical models each of the theoretical models has the finite amount of unknown parameters (moduli of elasticity e, characteristic times τc, viscosities η or the springpot parameters α). these parameters occur in analytical equations prescribing the behaviour of the models, e.g. the relaxation modulus. for our research the data obtained by experiments performed on laminated glass presented in [9] were used. the validation of the theoretical models was provided by fitting the parameters of the storage modulus. the python optimization tool from scipy was used for fitting the parameters, this tool uses the method of non-linear least squares to fit the analytical function to experimental data. the results are displayed in the following figures, where the blue dots illustrate the experimentally obtained data while the red curve is the fitted analytical solution. figure 7 shows the approximation by the 1-cell maxwell chain. however, the approximation is not very accurate. to improve the accuracy we need to add more maxwell cells to the generalized model. the increase of the number of maxwell cells from 1 to 3 is shown in figure 8. the approximation is good on the wider range of frequencies. as is shown in figure 9, the 30 vol. 40/2023 modelling of laminated glass interlayer by fractional viscoelasticity 10 5 10 3 10 1 101 103 angular velocity [rad/s] 106 107 st or ag e m od ul us [p a] fitted frac. m. chain experimental data figure 10. fitting experimental data of g′ to 1-cell fractional maxwell chain. 10 5 10 3 10 1 101 103 angular velocity [rad/s] 106 107 st or ag e m od ul us [p a] fitted frac. m. chain experimental data figure 11. fitting experimental data of g′ to 2-cell fractional maxwell chain. 7-cells maxwell chain approximates the experimental data well in the whole measured frequency range. on the other hand, the increase of the number of the cells from 1 to 7 also significantly increases the number of the unknown parameters that need to be fitted. the approximation by the fractional maxwell chain is shown in figures 10–12. even for the 1-cell fractional maxwell chain (figure 10) the accuracy of the approximation is significantly better compare to standard 1-cell maxwell chain (figure 7). however, the accuracy can also be improved by adding more fractional maxwell cells. figure 11 shows approximation by the 2-cells and figure 12 the 3-cells fractional maxwell chain. in the last mentioned case we can talk about almost perfect fit to the experimental data. the fractional maxwell chain shows one more significant advantage. while the curve of the standard maxwell chain become horizontal on the edges of the measured domain, the fractional maxwell chain still shows the growing trend of the curve. therefore, the fractional maxwell chain might be used for data extrapolation outside the experimentally measured domain. this section clearly shows the results that the frac10 5 10 3 10 1 101 103 angular velocity [rad/s] 106 107 st or ag e m od ul us [p a] fitted frac. m. chain experimental data figure 12. fitting experimental data of g′ to 3-cell fractional maxwell chain. tional maxwell chain can provide better approximations of the experimental data in comparison to standard generalized maxwell model. also it can be used for the data extrapolation for which the standard maxwell chain is not suitable. therefore the use of the fractional models might be beneficial despite their disadvantages which lie in the need of the more difficult mathematical apparatus. 3. conclusion fractional viscoelasticity seems to be an efficient tool to describe the behaviour of a polyvinyl butyral, the material of the interlayer of laminated glass. the principles of fractional calculus are introduced together with the springpot element. this rheological element can be used in connections with springs or dashpots to make more complex theoretical models. for our applications, the generalized maxwell chain models presented in its standard and fractional forms are sufficiently accurate and effective, particularly when considering the approximation of the experimental data. comparing both formulations, the fractional viscoelasticity requires a complex mathematical apparatus to calculate derivatives and integrals of non-integer order. on the other hand, by the use of fractional models the number of the fitted parameters may be reduced in compare to standard viscoelastic models. on top of that, it allows for extrapolation of data outside the experimentally measured domain. therefore, the fractional model can be directly exploited in fitting the experimentally measured master curve very accurately adopting a few parameters only. this smooth approximation can be in turn used in calibrating the maxwell chain model typically adopted in finite element simulations as implementation of the fractional model in the framework of the finite element method is still a subject of an ongoing research. 31 barbora hálková, jaroslav schmidt acta polytechnica ctu proceedings acknowledgements this publication was supported by the czech science foundation, the grant no. 22-15553s and by the grant agency of the czech technical university in prague, grant no. sgs22/031/ohk1/1t/11. references [1] w. laufs, a. luible. introduction on use of glass in modern buildings. tech. rep., epfl, laboratoire de la construction métallique icom, 2003. 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[2022-0815]. https://dspace.cvut.cz/handle/10467/102633 [8] r. gorenflo, f. mainardi, s. rogosin. mittag-leffler function: properties and applications. in volume 1 basic theory, pp. 269–296. de gruyter, berlin, boston, 2019. https://doi.org/10.1515/9783110571622-011 [9] t. hána, t. janda, j. schmidt, et al. experimental and numerical study of viscoelastic properties of polymeric interlayers used for laminated glass: determination of material parameters. materials 12(14):2241, 2019. https://doi.org/10.3390/ma12142241 32 https://doi.org/10.2749/sed010 https://doi.org/10.1115/1.3167616 https://doi.org/10.1063/1.4822863 https://dspace.cvut.cz/handle/10467/102633 https://doi.org/10.1515/9783110571622-011 https://doi.org/10.3390/ma12142241 acta polytechnica ctu proceedings 40:27–32, 2023 1 introduction 2 viscoelasticity 2.1 relaxation modulus of springpot 2.2 springpot under harmonic load 2.3 theoretical models 2.4 validation of theoretical models 3 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.34.0069 acta polytechnica ctu proceedings 34:69–74, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague environmental perspectives of fine grounded concrete powder: lca case study of light-weight concrete block jan peštaa, c, ∗, zdeněk prošekb, c a university of chemistry and technology, faculty of environmental technology, technická 5, 166 29 prague 6, czech republic b czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic c czech technical university in prague, university centre for energy efficient buildings, třinecká 1024, 273 43 buštěhrad, czech republic ∗ corresponding author: jan.pesta@cvut.cz abstract. the construction elements produced with low demand for primary resources are being developed in the building industry. however, not only consumption of primary resources but also environmental impacts of such products should be considered to increase the sustainability of the construction industry. to investigate the environmental impacts of lightweight concrete blocks, life cycle assessment (lca) was used as an analytical tool according to the en 15 804 + a2. firstly, the properties of fine grounded concrete powder (frwc) were measured to declare functionality and potential to use of the considered product. then, the lightweight concrete block containing frwc was designed. the potential environmental impacts of the lightweight concrete block were assessed for the following phases: raw material production, transport of materials, production of concrete block, and end of life of the block. the considered system boundaries includes the production of fine grounded concrete powder, which was produced by recycling demolished concrete structures. the results were related to 1 tun of lightweight concrete block. unsurprisingly, the calculated impact is mainly influenced by cement production. the total impact of the life cycle of 1 t of the lightweight block is 336 kg co2 eq. on the other hand, the lightweight block reached better results in the comparison with the aerated autoclaved concrete block in most of the categories. keywords: light-weight concrete, environmental aspect, lca. 1. introduction during the design phase of a product, the equilibrium among three pillars of sustainability (environmental, economic, and social) is balanced. the environmental perspective is based on the natural limitation of our planet, which can be considered as a system with limited resources. in such a system with limited resources, the production of new products is limited without developing of circularity of materials. the global circularity gap, the difference of materials consumed in the global economy and materials, which are recycled or reused, is more than 90 % [1]. especially, construction and demolition waste represent almost one-third of european waste [2]. moreover, the construction industry consumes more than 30 % of domestic material consumption in the czech republic [3]. to increase circularity, new construction products are being developed containing recycled content. this contribution describes the environmental performance of such a product, which was designed to use recycled concrete. one possible way, how to recycle concrete waste, is the production of fine grounded concrete powder. in comparison with processes, in which concrete is recycled to be used as recycled aggregate, this process due to fine grinding leads to the production of concrete powder with utilization as an admixture in the concrete mixture for lightweight concrete block. to assess environmental perspectives of fine grounded concrete powder (frwc), the case study considering the lightweight concrete block was conducted using life cycle assessment (lca) as a method to analyse environmental benefits and burdens caused by the life cycle of product or service. this method considers not only elementary flows of materials or elements and energies, but it also describes potential impacts, which can be caused by these flows in the environment. the life cycle assessment is standardized by iso 14 040:2006 [4], which describes conditions for the assessment of products and services. the results of such assessment can be used for environmental product declaration (epd) [5]. the epd for construction products is standardized by en 15 804+a2 [6], which also states the method for characterization of environmental impacts. in this contribution, lca was used to assess environmental impacts related to 1 t of the lightweight 69 https://doi.org/10.14311/app.2022.34.0069 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en jan pešta, zdeněk prošek acta polytechnica ctu proceedings concrete block. the assessed system boundaries include recycling of concrete and production of frwc. the environmental performance of considered concrete block was compared with environmental impacts declared in the epd of aerated autoclaved concrete. 2. materials and methods 2.1. lightweight concrete block the lca study was conducted to assess the environmental performance of a lightweight concrete block containing fine recycled waste concrete (frwc), which was used as a fine filler. the density of the block is 800 kg/m3. the proportions of the block are as follows: 500 mm length, 250 mm height, 175 widths. the concrete mixture is described in the following table 1. materials kg per t of blocks superplasticizer (spl) 10 cement i 42.5 r 331 frwc 490 recycled polypropylene fibres 8 water 157 foaming agent 7 table 1. description of materials used for 1 t of blocks. measured energy consumption in the process of preparing concrete mixture (mixing and foam preparation) was 15 kwh for the amount of concrete mixture used for the production of 1 t of blocks. the fire resistance and mechanical properties of the mixture were improved by using 100 % recycled polypropylene fibres. according to their manufacturer (trevos, s. r. o., czech republic), they have 32×10−3 mm in diameter and 4 mm in length. other properties are following: average tensile strength, <3.0 cn/dtex (∼272 mpa); average elongation, ≤50 %; density, 910 kg/m3; young modulus of elasticity, ∼4 gpa. the properties of concrete mixtures were improved by adding polycarboxylates as a superplasticizer (spl). 2.2. production of fine recycled waste concrete frwc was produced by crushing 100-years old concrete to a 0–32 mm fraction on the site of the recycling plant. as a by-side process, the steel scrap was removed from the crushed structure. in the following step, the crushed concrete was transported to the other recycling facility, where frwc was produced by grinding of 0-1 mm fraction of crushed concrete using a high-energy electric mill (sbd 800 from lavaris ltd.). the energy consumption of milling was estimated as 6.25 kw per milled tun of concrete. moreover, steel spins are considered to have a high rate of wear and so, the amount of steel spins used for the production of 1 t is calculated as 0.27 g per 1 t of recycled concrete. during the milling, no losses of recycled concrete are considered. after milling, frwc was transported to the plant, where concrete was prepared by mixing. 2.3. environmental assessment the environmental assessment was performed using the life cycle assessment method [7]. this method is standardized according to iso 14040 [4]. based on this standard, the study was conducted in the four basic steps: goal and scope definition, life cycle inventory, life cycle impact assessment, and life cycle interpretation. these four steps are iterative and the results of each of them can be used for assessment, which will be more appropriate for the goal and purpose of the study. the core category rules for building products are standardized by the european standard en 15 804 + a2 [6], which was applied in this study. in this study, the goal was to assess the environmental impacts of 1 t of concrete blocks containing frwc. the purpose of this study is to discuss the environmental burdens or benefits of using frwc for the production of concrete blocks with similar utilization as aerated autoclaved concrete. 2.3.1. system boundaries in compliance with en 15 804 + a2, the system boundaries include the following phase of the life cycle of the block: mining of raw materials, production of materials (including recycled concrete), transport of materials, preparing of concrete mixture, and end of life of the product (eol). the eol phase includes deconstruction, transport of waste materials, and landfilling of construction waste. during the life cycle, the use phase of the block is not considered. the reference service life is assumed to be higher than 50 years. 2.3.2. life cycle inventory gabi professional was used as software model to describe the boundaries of the system and to assess the functionality environmental impacts [8]. the model was used to describe elementary energies and materials flowing through considered system boundaries. in the following step, the potential environmental impacts caused by these elementary flows were assessed. for the creation of the model, specific data about the production of frwc were used. to model the production of other materials, generic data from the gabi database were used [9]. also, generic data were used for models of unspecific processes such as transport using trucks, czech electricity grid mix production, landfilling of construction waste, and diesel supply. 2.3.3. life cycle impact assessment in this step, elementary flows of described system are classified and characterized. classification of elementary flows represents the assigning of the flow to the impact category in the case, that the flow has the potential to cause impact in that category. one flow can be assigned in several impact categories. in the 70 vol. 34/2022 environmental perspectives of fine grounded concrete powder non renewable energy total raw materials transport production of end of life resources production concrete crude oil (resource) 2.12e+01 1.45e+01 1.31 3.24e-02 5.31 hard coal (resource) 7.06 5.77 8.42e-03 6.39e-01 6.45e-01 lignite (resource) 3.06e+01 2.42e+01 1.03e-02 5.70 6.60e-01 natural gas (resource) 9.94 8.22 1.01e-01 2.54e-01 1.36 peat (resource) 8.08e-03 6.95e-03 1.52e-04 7.88e-05 8.98e-04 uranium (resource) 3.41e-04 2.45e-04 4.97e-07 8.52e-05 1.05e-05 table 2. consumption of non renewable energy resources during life phases of 1 t of lightweight concrete block (consumption in kg). characterization, the potential impact of flow is calculated. the calculation is performed by multiplying of the amount of a flow by its characterization factor for each environmental indicator. characterization factors are estimated by characterization models and in this study, characterization models were used in compliance with the en 15 804 + a2. the calculated results of some environmental indicators represent the potential environmental impact in the category. 3. results results of environmental assessment of lightweight concrete block containing fine recycled concrete powder are described in the following chapter. 3.1. life cycle inventory outputs the life cycle of the lightweight concrete block can be assessed based on the consumption of resources. in table 2, the energy resource consumption is described. regarding consumption of crude oil, the processes related to transport in the life cycle are connected with lower consumption than raw material production. moreover, natural gas consumption is significantly influenced by the end of life processes. however, the processes of raw materials production have the highest consumption regarding all of the considered energy resources. 3.2. potential environmental impacts in this study, environmental indicators according to en 15 804 + a2 were applied to assess potential environmental impacts related to 1 t of the lightweight block. the results of this assessment are described in table 3. the raw materials production phase is responsible for the biggest contribution to impacts in each impact category. however, the eol phase has also a high impact in the category of water use (1.70 m3 world equiv.), which represents almost 21 % of the whole impact in this category. similarly, the production of concrete contributes almost 20 % to the impact in the ionising radiation category. this result is mainly influenced by the production of electricity for processes. nevertheless, the production phase and transport processes cause negligible impact compared to the raw production phase. 4. discussion 4.1. contribution analysis the contribution of each process was analysed to declare the processes with the highest influence on potential impact assessed for the life cycle of 1 t of the lightweight block. in table 4, the relative contribution of processes to impact in a category are declared. categories climate change – fossil, biogenic, and land-use change are not included. in the table, processes contributing more than 10 % of the total impact in each category are stated. results of contribution analyses were conducted only with environmental indicators, which are characterized according to en 15 804 + a2. the concrete recycling process contributes beneficially to several impact categories. unsurprisingly, recycling of concrete decreases the overall impact in the resource use (mineral and metals) category. moreover, the recycling of concrete cause a beneficial impact in the water use category. on the other hand, the process, which is most contributing to several categories, is cement production. the impact of this process represents more than 80 % of impact in the climate change (total) category. also, it is significantly contributing to the impact in categories such as acidification, eutrophication, photochemical ozone creation, resource use (fossil), and water use. the impact in these categories is also influenced by the contribution of landfilling process, which describes environmental impacts caused by construction inert material disposed on the landfill. superplasticizer production is another process, which affects the impacts of considered lightweight concrete blocks. this process, which represents the production of 10 kg of superplasticizers, contributes 96 % to ozone depletion and it significantly increases the results of three other impact indicators (eutrophication freshwater; resource use, mineral and metals; water use). 4.2. comparison of the lightweight block with an aerated autoclaved concrete block the lightweight concrete block is intended to be a commercial product and so its environmental performance was compared to another block, which is available 71 jan pešta, zdeněk prošek acta polytechnica ctu proceedings total raw materials transport production of end of production concrete life climate change total [kg co2 eq.] 3.36e+02 2.99e+02 4.47 9.76 2.31e+01 climate change, fossil [kg co2 eq.] 3.36e+02 2.98e+02 4.44 9.68 2.35e+01 climate change, biogenic [kg co2 eq.] 2.23e-01 6.44e-01 -5.29e-03 6.10e-02 -4.76e-01 climate change, land use and land use change [kg co2 eq.] 3.81e-01 2.19e-01 3.63e-02 1.30e-02 1.13e-01 ozone depletion [kg cfc-11 eq.] 1.79e-06 1.79e-06 8.77e-16 1.19e-13 6.10e-14 acidification [mole of h+ eq.] 7.05e-01 5.13e-01 1.47e-02 2.25e-02 1.55e-01 eutrophication, freshwater [kg p eq.] 3.70e-03 3.61e-03 1.32e-05 2.40e-05 5.05e-05 eutrophication, marine [kg n eq.] 1.94e-01 1.32e-01 6.76e-03 4.45e-03 5.09e-02 eutrophication, terrestrial [mole of n eq.] 2.11 1.43 7.55e-02 4.64e-02 5.62e-01 photochemical ozone formation, human health [kg nmvoc eq.] 5.75e-01 4.19e-01 1.33e-02 1.23e-02 1.30e-01 resource use, mineral and metals [kg sb eq.] 2.35e-04 2.31e-04 3.94e-07 1.45e-06 2.18e-06 resource use, fossils [mj] 2.14e+03 1.62e+03 5.91e+01 1.44e+02 3.13e+02 water use [m3 world equiv.] 9.66 7.78 4.12e-02 1.36e-01 1.70 resource use indicators use of renewable primary energy (pere) [mj] 3.07e+02 2.39e+02 3.40 3.14e+01 3.36e+01 total use of renewable primary energy resources (pert) [mj] 3.07e+02 2.39e+02 3.40 3.14e+01 3.36e+01 use of non-renewable primary energy (penre) [mj] 2.14e+03 1.62e+03 5.93e+01 1.44e+02 3.13e+02 total use of non-renewable primary energy resources (penrt) [mj] 2.14e+03 1.62e+03 5.93e+01 1.44e+02 3.13e+02 use of renewable secondary fuels (rsf) [mj] 1.64e-22 1.64e-22 0.00 0.00 0.00 use of non renewable secondary fuels (nrsf) [mj] 1.93e-21 1.93e-21 0.00 0.00 0.00 use of net fresh water (fw) [m3] 6.34e-01 5.27e-01 3.90e-03 4.64e-02 5.68e-02 output flows and waste categories hazardous waste disposed (hwd) [kg] 8.40e-04 8.40e-04 3.13e-09 3.16e-08 2.73e-08 non-hazardous waste disposed (nhwd) [kg] 1.00e+03 2.78 9.31e-03 5.75e-02 1.00e+03 radioactive waste disposed (rwd) [kg] 6.43e-02 4.37e-02 1.08e-04 1.82e-02 2.28e-03 optional indicators particulate matter [disease incidences] 8.46e-06 6.67e-06 8.51e-08 1.71e-07 1.54e-06 ionising radiation, human health [kbq u235 eq.] 7.52 6.05 1.57e-02 1.21 2.43e-01 ecotoxicity, freshwater [ctue] 1.10e+03 8.09e+02 4.39e+01 5.28e+01 1.97e+02 human toxicity, cancer [ctuh] 7.65e-08 5.60e-08 8.87e-10 1.10e-09 1.85e-08 human toxicity, non-cancer [ctuh] 5.00e-06 2.92e-06 5.24e-08 7.43e-08 1.95e-06 land use [pt] 4.51e02 3.03e+02 2.03e+01 4.74e+01 8.09e+01 table 3. results of environmental indicators for the life cycle of 1 t of the lightweight block according to en 15 804 + a2. 72 vol. 34/2022 environmental perspectives of fine grounded concrete powder concrete cement landfilling of polypropylene steel spl recycling (cem i 42.5) construction fibres parts process production matter production production climate change total -2.83 80.94 4.39 5.31 0.31 3.30 ozone depletion 0.00 0.00 0.00 0.00 3.01 96.04 acidification -1.32 61.45 15.28 3.33 0.76 5.52 eutrophication, freshwater 0.13 3.43 0.69 0.87 11.84 80.51 eutrophication, marine 1.28 55.09 14.41 3.60 0.62 3.74 eutrophication, terrestrial 1.50 55.12 14.53 3.48 0.52 3.54 photochemical ozone formation, human health -0.08 60.15 14.74 4.73 0.52 4.84 resource use, mineral and metals -10.97 7.70 0.61 1.63 9.91 89.69 resource use, fossils -3.35 33.49 9.38 28.63 0.66 11.14 water use -24.03 28.51 16.77 1.50 5.14 58.59 table 4. the relative contribution of processes (only processes with relative contribution higher than 10 % of total impact in a category are stated), %. frwc block aac climate change total [kg co2 eq.] 2.51e+02 2.32e+02 ozone depletion [kg cfc-11 eq.] 1.43e-06 1.90e-03 acidification [mole of h+ eq.] 4.40e-01 6.28e-01 eutrophication, freshwater [kg p eq.] 2.92e-03 2.35e-02 photochemical ozone formation, human health [kg nmvoc eq.] 3.56e-01 1.03e+00 resource use, mineral and metals [kg sb eq.] 1.86e-04 2.69e-03 resource use, fossils [mj] 1.46e+03 1.79e+03 water use [m3 world equiv.] 6.37e+00 1.40e+01 table 5. comparison of the environmental performance of lightweight block with aac block (results of a1-a3 modules. on the market. to investigate products with similar properties and their environmental performance, the environdec database was used as a database of products [10], which were assessed using lca and their environmental performance was declared in environmental product declaration according to en 15 804 + a2. to ensure comparability of products based on their functionality, the aerated autoclaved concrete (aac) blocks were searched. the aac produced by gasbeton was selected as the most appropriate alternative [11]. the selected product is called sysmic idro (density 580 kg/m3). the comparison of their environmental performance related to the 1 m3 od product is presented in table 5. in most of the categories, the lightweight block reached better results than the aac block. on the other hand, the aac block contributes less to the impact in the category climate change. this comparison is limited by system boundaries. the environmental performance of the aac block was calculated only in the cradle-to-gate scope, which includes only a1-a3 phases, so the eol phase is not included in this comparison. the result of the frwc block was calculated based on a measured density of 800 kg/m3. 5. conclusion the lca was performed for the lightweight concrete block. the environmental indicators were used according to en 15 804 + a2 and so the results can be compared with another construction product with 73 jan pešta, zdeněk prošek acta polytechnica ctu proceedings the same function. in the comparison with aac, the lightweight block reached lower environmental impacts in most of the categories. on the other hand, the production of lightweight block caused a higher impact (251 kg co2 eq.) than the production of aac (232 co2 eq.) considering the climate change indicator. unsurprisingly, the cement production influence mainly the result of the climate change (total) indicator. however, the recycling of concrete producing frwc have a beneficial impact not only in this category (-9.41 kg co2 eq.) but also in the categories such as resource use, mineral and metals ( 2.58e-05 kg sb eq.) and water use (-2.32 m3 world eq.). therefore, further research will be conducted to optimize cement reduction and increase the use of frwc. acknowledgements this work was supported by the grant of specific university research – grant no a1_ftop_2021_003. references [1] the circularity gap report 2021. [2] strategy for a sustainable built environment. [3] zpráva o životním prostředí české republiky 2015. tech. rep., 2015. [4] m. finkbeiner, a. inaba, r. tan, et al. the new international standards for life cycle assessment: iso 14040 and iso 14044. the international journal of life cycle assessment 11(2):80–85, 2006. [5] iso. environmental labels and declarations — type iii environmental declarations — principles and procedures(iso 14025:2006), 2006. [6] cen. en 15804:2012+a2:2019 sustainability of construction works environmental product declarations core rules for the product category of construction products. [7] j. guinee. handbook on life cycle assessment operational guide to the iso standards. international journal of life cycle assessment 6(5):255–255, 2001. https://doi.org/10.1007/bf02978784. [8] life cycle assessment lca software: gabi software. [9] t. kupfer, c. m. colodel, m. kokborg, et al. gabi database and modelling principles, 2020. [10] environdec, the international epd system. [11] gasbeton, environmental product declaration. 74 https://doi.org/10.1007/bf02978784 acta polytechnica ctu proceedings 34:69–74, 2022 1 introduction 2 materials and methods 2.1 lightweight concrete block 2.2 production of fine recycled waste concrete 2.3 environmental assessment 2.3.1 system boundaries 2.3.2 life cycle inventory 2.3.3 life cycle impact assessment 3 results 3.1 life cycle inventory outputs 3.2 potential environmental impacts 4 discussion 4.1 contribution analysis 4.2 comparison of the lightweight block with an aerated autoclaved concrete block 5 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0117 acta polytechnica ctu proceedings 40:117–122, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague return mapping scheme for the hoek-brown model with tension cut-off tereza žalská∗, michal šejnoha czech technical university in prague, faculty of civil engineering, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: tereza.zalska@fsv.cvut.cz abstract. the present paper revisits a stress update procedure for the hoek-brown plasticity model enhanced by a rankine type of tension cut-off failure criterion. limiting tensile stresses not only supports the behavior of weak rock masses with a very low tensile strength but in combination with the original hoek-brown model improves robustness of the stress update procedure, i.e. the stress return mapping algorithm. herein, the primary focus is on the stress return from a space of inadmissible trial stresses for which neither the hoek-brown yield surface nor its derivative can be evaluated. all potential scenarios are thoroughly discussed both in the framework of singleand multi-surface plasticity. the presented procedures were implemented and verified with the help of the geo5 fem software. keywords: hoek-brown criterion, rankine criterion, nonlinear yield function, rock mechanics, finite element method, tension cut-off. 1. introduction in geological engineering, shear strength properties of the rock mass are mostly expressed by the parameters entering the formulation of the hoek-brown (hb) plasticity model. these can be transformed into the pair of basic shear strength parameters, i.e, cohesion c and friction angle φ, exploiting similarity between the hoek-brown and mohr-coulomb (mc) yield functions. then the numerical analysis can be simply governed by the mc model. however, formulation of the hb model is based on the assumption of a nonlinear development of the rock mass strength with increasing loading. on the other hand, equivalent shear strength parameters derived by hoek and brown in [1] are independent of the current stress state. therefore, this approach requires a proper estimation of the stress range for obtaining comparable results, see [2]. in light of this, a direct application of the hb model appears more suitable, which also promoted a recent implementation of the hb model into geo5 fem software [3]. in a numerical solution, violation of the yield function is followed by a correction step bringing the stresses back to the yield surface. formulation of a suitable stress return algorithm typically calls for the determination of the first derivative of a given yield function. point out that the hb yield function is not defined for the maximum trial principal stress exceeding the tensile strength of the rock mass making the standard return mapping inapplicable. to overcome this obstacle, the present paper introduces a stress update algorithm that incorporates the rankine failure criterion in conjunction with the hb model when moving from inadmissible to admissible stress space. within this concept, the rankine failure criterion may also represent a less or more significant reduction of the tensile strength. 2. governing equations application of the generalized hoek-brown criterion covers the whole range of rock masses from an intact one to a highly disturbed weak rock mass. the input parameters of the hb model are [4]: • uniaxial compressive strength σci and hoek-brown constant mi provided by series of triaxial testing, • geological strength index (gsi) used as a classification number of the geological quality of rock mass estimated from in situ geological observation, • disturbance factor d describing the rate of disturbance by prior excavations or blasting, also estimated from in situ geological observation. the material constants mb, s and a govern the shape of the yield surface and are given by empirical equations as mb = mi exp gsi − 100 28 − 14d , (1) s = exp gsi − 100 9 − 3d , (2) a = 1 2 + 1 6 ( exp −gsi 15 − exp −20 3 ) . (3) 2.1. hoek-brown yield function assuming a standard order of principle stresses σ1 > σ2 > σ3, the generalized yield function is defined as f13(σ1,σ3) = σ1 − σ3 − σci ( s − mb σ1 σci )a , (4) where σ1 and σ3 are the maximal and minimal effective principal stresses, respectively. unlike typical geotechnical notation, the compressive stresses are assumed negative and σ1 has a meaning of confining 117 https://doi.org/10.14311/app.2023.40.0117 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en tereza žalská, michal šejnoha acta polytechnica ctu proceedings pressure. equation (4) represents the main sector of the yield surface in the principle stress space. for specific cases of triaxial compression (lode’s angle θ = 30◦) and triaxial extension (θ = −30◦) two more sectors need to be considered. these are defined as f23(σ2,σ3) = σ2 − σ3 − σci ( s − mb σ2 σci )a , (5) f12(σ1,σ2) = σ1 − σ2 − σci ( s − mb σ1 σci )a . (6) in the principal stress space, the hb model plots as a irregular hexagonal pyramid with curved edges and surfaces, see figure 1. assuming the uniaxial tensile strength be equal to the biaxial one (σ3 = σ1 = σt) yields the tensile strength of the rock mass adopting equation (4) as σt = sσci mb . (7) for the intact rock it holds gsi = 100 which gives a = 0.5. as the minimum value of gsi is equal to 0, the exponent a ranges from 0.5 to approximately 0.666. it is obvious that setting σmax > σt generates singularity in eqautions (4)–(6). the yield condition is not defined beyond the apex and hence it cannot be evaluated. σm σ2 σ3 σ1 figure 1. hoek-brown yield surface plotted in principal stress space. expressing the yield function in terms of stress invariants, i.e., the mean stress σm, the deviatoric stress measure j and the lode’s angle θ gives fhb (σm,j,θ) = (2j cos θ) 1/a − sσ1/aci + + ( cos θ − sin θ √ 3 ) mbjσ 1/a−1 ci + mbσmσ 1/a−1 ci . (8) note that formulation (8) is not suitable for the complex return mapping procedure, as the derivative of the yield surface along the triaxial compression and triaxial extension edges (lode’s angle θ = ±30◦, respectively) is singular, see [5] with reference to the mc model. 2.2. plastic potential plastic potential functions are formulated on the basis of the mc yield criterion as g13(σ1,σ3) = σ1 1 + sin ψ 1 − sin ψ − σ3, (9) g23(σ2,σ3) = σ2 1 + sin ψ 1 − sin ψ − σ3, (10) g12(σ1,σ2) = σ1 1 + sin ψ 1 − sin ψ − σ2, (11) where ψ is well known dilation angle. constant rate of dilation is assumed in the present formulation. vectors normal to the plastic potential surfaces governing the direction of the plastic yielding are therefore independent of the current stresses and take the form n1g = { 1 + sin ψ 1 − sin ψ ; 0; −1 } t, (12) n2g = { 0; 1 + sin ψ 1 − sin ψ ; −1 } t, (13) n3g = { 1 + sin ψ 1 − sin ψ ; −1; 0 } t. (14) more advanced formulations of the potential functions are also available in literature, see, e.g., carranzatorres and fairhurst [6] who proposed the plastic potential function based on the mobilized value of the dilation angle ψmob depending on current stress state. however, such formulations go beyond the present scope. 2.3. rankine yield function yield surfaces of the rankine type are often used as the additional tension cut-off surfaces. the model assumes the material to be isotropic and unlike the hb model, the stress return mapping is in general driven by an associated flow rule. the rankine yield surfaces are defined as t1(σ1,σt) = σ1 − σt, (15) t2(σ2,σt) = σ2 − σt, (16) t3(σ3,σt) = σ3 − σt, (17) where σt ≤ σt is the prescribed tensile strength. the vectors normal to individual sectors of the yield surface are for the assumed associated flow rule given by n1t = {1; 0; 0} t, (18) n2t = {0; 1; 0} t, (19) n3t = {0; 0; 1} t. (20) note that apart from a physically accepted reduction of a tensile stress σt, recall equation (7), the return to the rankine yield surface will be adopted as an intermediate step when an attempt to return to the hb yield function is made from an inadmissible trial 118 vol. 40/2023 hoek-brown plasticity model with tension cut-off stress space1. to avoid failing the return process we consider an upper limit of the tensile strength σmaxt = σt − σ∗ where σ∗ = 1 kpa is assumed in geo5 fem. for the sake of completeness we present the rankine yield function, which plots as a regular triangular pyramid in the principal stress space, also in terms of stress invariants as fr(σm,j,θ) = 2 √ 3j cos ( θ + π 6 ) + + 3σm − 3σt. (21) 3. stress update algorithm in the presented approach, the rock mass is treated as isotropic and no hardening is considered. stress update algorithm respects the standard theory of plasticity. elastic region of the strain is bounded by the yield surface. in case of plastic yielding both the yield and consistency condition must be satisfied. the geo5 fem software assumes the stress-strain relation to be linearly elastic perfectly plastic for the hb model. the total strain increment decomposes into elastic and plastic parts as dε = dεel + dεpl. (22) the increment of stress relates to the elastic part of the total strain and follows the linear relation dσ = del( dε − dεpl). (23) for non-associated plasticity the plastic part of the strain increment is defined as dεpl = ∆λ ∂g(σ) ∂σ , (24) with ∆λ being positive for plastic yielding, otherwise it is equal to 0. an associated plastic flow rule assumes g(σ) = f(σ). as already mentioned, a non-associated flow rule is applied when returning to the hb yield surface, while associated plasticity is adopted for the rankine yield surface. derivation of the plastic strain increment follows standard two step procedure employing elastic predictor and plastic corrector steps. the former step gives the trial stresses in the form σtr = σj−1 + del dε. (25) once the yield condition is violated, i.e., f(σtr ) > 0, the return mapping is carried out to bring the trial stresses back to the yield surface. depending on the orientation of the elastic guess and the direction of the plastic strain increment, reordering of the principal stresses may occur, i.e. σ1 > σ2 > σ3 may no longer apply. as will be graphically shown in the next section, reordering of principal stresses requires formulation 1the trial stress space is assumed inadmissible whenever condition σmax > σt occurs regardless of whether or not the hb model is defined. of the return mapping algorithm in the framework of multi-surface plasticity. this is because the updated stress, although theoretically located on the yield surface, gives lode’s angle which falls outside the limits −30◦ ≤ θ ≤ 30◦. in such a case, the correct stress lies on the triaxial edge, where two sectors of the yield surface intersect so both corresponding surfaces become active in return mapping. 3.1. single yield surface return strategy implicit return mapping is applied using the newtonraphson method in the form (∆λ)i+1 = (∆λ)i − (f13)i ( df13)i , (26) where df13 is the derivative of the yield function associated with the main sector and ∆λ is the plastic strain increment. the updated stress is provided by σi+1 = σtr − (∆λ)i+1deln1g, (27) where the direction of the plastic strain increment n1g remains constant during the return mapping, recall equations (12)–(14). writing the direction of principal plastic corrector as deln1g = σ pl = { σ pl 1 ,σ pl 2 ,σ pl 3 } t (28) gives σ1(∆λ) = σtr1 − ∆λσ pl 1 , (29) σ2(∆λ) = σtr2 − ∆λσ pl 2 , (30) σ3(∆λ) = σtr3 − ∆λσ pl 3 . (31) substituting equations (29)–(31) into equation (4) gives the yield function f13 in the form f13(∆λ) = (σtr1 − σ tr 3 ) − ∆λ(σ pl 1 − σ pl 3 )− − σci [ s − mb σci ( σtr1 − ∆λσ pl 1 )]a (32) and the corresponding derivative w.r.t. ∆λ as df13(∆λ) = σpl3 − σ pl 1 − − ambσ pl 1 [ s − mb σci ( σtr1 − ∆λσ pl 1 )]a−1 . (33) return to the rankine yield surface t1 follows the same minimization procedure. owing to the linear form of the yield function, the correct increment of the plastic strain is found in one iteration step. therefore ∆λ = − t1 dt1 . (34) the tensile yield surface defined in terms of the plastic multiplier is given by t1 = σtr1 − ∆λσ pl 1 − σt (35) 119 tereza žalská, michal šejnoha acta polytechnica ctu proceedings σ1 σ2 σ3 0 50 100 150 200 0 20 40 60 80 100 f hb f r f hbf r -σm (kpa) j ( k p a ) f r f hb -σ1 (kpa) -σ 3 ( k p a ) σm figure 2. non-associated return to the hb surface from trial stress violating t1 in two consecutive steps (ψ = 0): first from trial stress (red point) onto the rankine yield surface (orange point), second from rankine surface to hb surface (green point) plotted in the deviatoric plane (left), σm − j plane (middle), σ1 − σ3 plane (right). and the derivative w.r.t. ∆λ is therefore constant dt1 = −σpl1 . (36) returning from the trial stress space violating the rankine failure criterion considers two scenarios: • correct return to the hb yield surface is expected: this is a two-step return where returning to the rankine yield surface is performed first assuming the direction of plastic corrector given by the hb model, i.e., σpl = deln1g (non-associated flow rule). the procedure then continues with the second step bringing the stresses back onto the hb yield surface along the same direction. the total plastic strain increment is then the sum of plastic strain increments generated by both steps. see figure 2. • correct return to the rankine yield surface is expected: this is a one-step return along the direction given by σpl = deln1t (associated flow rule). 3.2. triaxial lines return strategy two active sectors of the hb yield surface apply to both the state of triaxial compression and triaxial extension. considering triaxial stress state in compression for the determination of the multi-surface plasticity return, two yield functions, f13 and f23, are active and the iterative scheme takes the form{ ∆λ1 ∆λ2 }i+1 = { ∆λ1 ∆λ2 }i −(h−1)i { f13 f23 }i , (37) where the jacobian matrix is composed of the partial derivatives of the yield functions with respect to two plastic multipliers ∆λ1 and ∆λ2. therefore h =   ∂f13 ∂(∆λ1) ∂f13 ∂(∆λ2) ∂f23 ∂(∆λ1) ∂f23 ∂(∆λ2)   . (38) the vector of updated principle stresses for the current iterative step is given by σi+1 = σtr − (∆λ1)i+1deln1g − (∆λ2) i+1deln2g. (39) vectors providing the orientation of the plastic corrector are defined as σ pl 1 = d eln1g = { σ pl 11; σ pl 12; σ pl 13 }t , (40) σ pl 2 = d eln2g = { σ pl 21; σ pl 22; σ pl 23 }t (41) and individual components of the stress vector for j = 1 → 3 σj (∆λ1, ∆λ2) = σtrj − ∆λ1σ pl 1j − ∆λ2σ pl 2j. (42) expressing the yield function in terms of two plastic multipliers provides the yield function f13 as f13(∆λ1, ∆λ2) = = (σtr1 − σ tr 3 ) − ∆λ1(σ pl 11 − σ pl 13) − ∆λ2(σ pl 21 − σ pl 23)− − σci [ s − mb σci ( σtr1 − ∆λ1σ pl 11 − ∆λ2σ pl 21 )]a (43) and its partial derivatives in the form h11 = σ pl 13 − σ pl 11 − ambσ pl 11[ s − mb σci ( σtr1 − ∆λ1σ pl 11 − ∆λ2σ pl 21 )]a−1 , (44) h12 = σ pl 23 − σ pl 21 − ambσ pl 21[ s − mb σci ( σtr1 − ∆λ1σ pl 11 − ∆λ2σ pl 21 )]a−1 . (45) the other yield function and components of the h matrix would be expressed similarly and the same approach is applied to the state of triaxial extension with sectors f13 and f12 being simultaneously active. attention should be also paid to the triaxial corner of the tensile yield surface. in this case, the proposed approach is adopted for the intersection of the yield functions t1 and t2 and, similarly to a single surface plasticity with one iteration step needed. again, both non-associated and associated plasticity is considered in return mapping scheme depending on whether the stress is expected to be brought back to hb model or directly to the rankine yield surface, recall the last paragraph in the previous section. as an example of the former case, see figure 3. 120 vol. 40/2023 hoek-brown plasticity model with tension cut-off figure 3. demonstration of the return to the triaxial corners of the hb yield surface from trial stress violating tensile yield condition (ψ = 0 is assumed): σtr violates t1 condition (tc corner) σtr ′ violates both t1, t2 conditions (tc corner) σtr∗ violates t1 condition (te corner). 4. complex return mapping scheme here, only the return from an inadmissible trial stress space is discussed. as already explained in the previous sections, a direct return of the trial stress for the condition σtr1 > σt onto the hb yield surface fails as the hb yield function is undefined. the return mapping therefore begins with evaluating the tensile condition (15) employing the reduced tensile strength σt ≤ σt − σ∗. violating this condition calls a non-associated return onto the rankine yield function controlled by the plastic potential formulated for the hb model. if the stress point violates the hb criterion, a non-associated return from the point on the rankine surface onto the hb yield surface follows, see figure 2. otherwise, associated plasticity is adopted for bringing the original trial stress back onto the rankine tensile yield surface, recall the last paragraph in section 3.1. special attention should be paid to the intersection of t1 and f13 surfaces with reference to figure 5. the proposed algorithm first predicts that the return mapping will continue to the hb surface after being preceded by a non-associated return to the tensile yield surface. if no violation of the hb criterion is detected after the first step, the associated return of the trial stress to the tensile surface is called. after this attempt, violation of the hb condition may eventually occur. if it does, the hb criterion is substituted with the linear function fsub, see equation (46), which satisfies t1 condition (σ1 = σt), and the trial stress is shifted to the intersection of two planes using multi-surface plasticity. fsub(σ3) = σt − σ3 − σci ( s − mb σt σci )a (46) it is appropriate to point out that reordering of the principle stresses, and thus detection of the triaxial corner return, may occur during the described steps. in such a case, the above two-step procedure would also be performed in the framework of multi-surface plasticity. a graphical representation is evident in figure 3. the complete return mapping scheme including all potential scenarios is then summarized in figure 4. figure 4. full return mapping scheme including all potential scenarios. 121 tereza žalská, michal šejnoha acta polytechnica ctu proceedings σ1 σ2 σ3 σ1 σ2 σ3 σ1 σ2 σ3 10 0 10 20 0 5 10 15 20 25 10 0 10 20 0 5 10 15 20 25 10 0 10 20 30 0 5 10 15 20 25 f hb f r -σm (kpa) j ( k p a ) j ( k p a ) -σm (kpa) -σm (kpa) j ( k p a ) delng1 d eln t 1 figure 5. sequence of the return attempts leading to correct stress lying on the intersection of t1 and f13 (ψ = 0) plotted in the deviatoric planes (top) and σm − j planes (bottom). 5. conclusions a detailed description of the return mapping algorithm for the case of inadmissible trial stresses, here defined as the stress space for which σtr1 > σt, is presented. to the authors knowledge the number of contributions addressing this issue is limited. in [7] a smooth variant of the hb model is presented and the tension cut-off condition is introduced via an allowable tensile mean stress instead of actual tensile strength. in [8, 9] the authors also deal with the issue of tensile stresses. however, the approach is not explained in all details and its effectivity is not obvious. the present approach exploits simplicity of the rankine yield surface, which is often used as a tension cut-off criterion. the return of inadmissible trial stresses to the main sector or triaxial edges of the hb yield surface assumes two consecutive steps. first, the stress is brought back to the rankine yield surface defined by the reduced tensile strength. second, the implicit iterative procedure for shifting the stress onto the hb surface is applied. both steps assume the same orientation of the plastic corrector. it is shown, that the prediction of all possible situations, which can lead to exceeding the tensile strength and failing the solution, results in the complex algorithm further complicated by the assumption of associated plasticity for tensile yielding. this assumption becomes active once the two-step return procedure onto the hb surface fails and the trial stress is then brought back directly to the rankine yield surface. this algorithm was implemented into geo5 fem software and thoroughly tested. acknowledgements the support provided by the sgs project no. sgs22/030/ohk1/1t/11 and by the gačr project no. 22-12178s is gratefully acknowledged. references [1] e. hoek, c. carranza-torres, b. corkum. hoek-brown failure criterion – 2002 edition. in proceedings of the 5th north american rock mechanics symposium – narms-tac, pp. 267–273. 2002. [2] t. poklopová, m. šejnoha. comparing the hoek-brown and mohr-coulomb failure criteria in fem analysis. in acta polytechnica ctu proceedings, vol. 30, pp. 69–75. 2021. https://doi.org/10.14311/app.2021.30.0069 [3] fine-ltd. geo5 fem, 2022. [2022-09-13]. http://www.fine.cz/ [4] e. hoek, e. t. brown. practical estimates of rock mass strength. international journal of rock mechanics and mining sciences 34(8):1165–1186, 1997. https://doi.org/10.1016/s1365-1609(97)80069-x [5] j. huang, d. v. griffiths. observations on return mapping algorithms for piecewise linear yield criteria. international journal of geomechanics 8(4):253–265, 2008. https: //doi.org/10.1061/(asce)1532-3641(2008)8:4(253) [6] c. carranza-torres, c. fairhurst. the elasto-plastic response of underground excavations in rock masses that satisfy the hoek-brown failure criterion. international journal of rock mechanics and mining sciences 36(6):777–809, 1999. https://doi.org/10.1016/s0148-9062(99)00047-9 [7] r. g. wan. implicit integration algorithm for hoek-brown elastic-plastic model. computers and geotechnics 14(3):149–177, 1992. https://doi.org/10.1016/0266-352x(92)90031-n [8] j. clausen. efficient non-linear finite element implementation of elasto-plasticity for geotechnical problems. ph.d. thesis, aalborg universitet, 2007. [9] j. clausen, l. damkilde. an exact implementation of the hoek-brown criterion for elasto-plastic finite element calculations. international journal of rock mechanics and mining sciences 45(6):831–847, 2008. https://doi.org/10.1016/j.ijrmms.2007.10.004 122 https://doi.org/10.14311/app.2021.30.0069 http://www.fine.cz/ https://doi.org/10.1016/s1365-1609(97)80069-x https://doi.org/10.1061/(asce)1532-3641(2008)8:4(253) https://doi.org/10.1061/(asce)1532-3641(2008)8:4(253) https://doi.org/10.1016/s0148-9062(99)00047-9 https://doi.org/10.1016/0266-352x(92)90031-n https://doi.org/10.1016/j.ijrmms.2007.10.004 acta polytechnica ctu proceedings 40:117–122, 2023 1 introduction 2 governing equations 2.1 hoek-brown yield function 2.2 plastic potential 2.3 rankine yield function 3 stress update algorithm 3.1 single yield surface return strategy 3.2 triaxial lines return strategy 4 complex return mapping scheme 5 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.35.0008 acta polytechnica ctu proceedings 35:8–13, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague digital track map for the vexa expert system adam hlubuček czech technical university in prague, faculty of transportation sciences, department of transport telematics, konviktská 20, 110 00 prague 1, czech republic, prague, czech republic correspondence: hlubuada@fd.cvut.cz abstract. this paper aims to summarize possible solutions of a digital track map proposal, designed to become a data source for the vexa decision-making system. the vexa project is carried out in cooperation of the department of applied mathematics at the czech technical university in prague – faculty of transportation sciences and the ažd praha s. r. o. company. the digital track map is an indispensable data source describing the railway line intended for the vexa supervised trains operation and its surroundings. being created in the arcgis pro software environment, the digital map allows individual objects to be visualized and handled as standard gis features and described by related data stored in a geodatabase. depending on the object class, the object location can be expressed in different ways. not only physical location (directly expressible using gis features) but also functional location (as the objects relate to a particular track) must be considered. keywords: digital track map, vexa, autonomous railway vehicle, expert system, gis. 1. introduction the effort to reduce the influence of the critical human factor in the railway sector leads to the development of numerous process information systems. such systems include, for example, automatic train control systems, automatic train operation systems, various forms of railway traffic management systems and decision-making systems. whereas many of these systems require accurate infrastructure data, the need for thorough description of the railway infrastructure is growing. this is not different in the case of the expert system that is being developed within vexa. the vexa research project is carried out in cooperation of the department of applied mathematics at the czech technical university in prague – faculty of transportation sciences and the ažd praha s.r.o. control and signalling systems supplier. the vexa expert system itself consists of several interconnected modules providing various functionalities leading to perform the tasks of an autonomous railway vehicle. based on the evaluation of defined inputs, it aims to substitute decision-making processes of a train driver. these inputs can be outputs from detectors (e. g. obstacle, fire) and status reporting systems (as regards the individual modules) and track condition monitoring. with the use of a knowledge base and decision rules (railway operation rules, decision algorithms taking into account the driver’s behaviour), vexa is supposed to evaluate the conditions for running a railway vehicle in order to grant a permit to start moving or to issue a request to stop the train. with the use of the vexa system, it is expected to reach the grade automation of the level goa3 or goa4 (i. e. driverless or unattended train operation) [1, 2]. in order to provide the vexa system with information concerning the environment in which the train is running, the digital track map is being developed. the map is supposed to describe the railway line intended for the vexa supervised trains operation as well as its surroundings. the digital track map data represent one of the important inputs for the vexa system decision making processes. among others, the data will be compared with detected scene in order to recognise unwanted objects. 2. logical architecture of the vexa system the onboard vexa unit is supposed to consist of four following modules: • perception input module • map matching module • incident analysis module • incident solver module [1] the perception input module is responsible for aggregating inputs from train detectors of different types. as the set of sensors may differ for individual implementations of an autonomous train, the module is being designed to be configurable. the map matching module is in charge of managing information about the known surroundings of the autonomous train. based on the unit location information, it provides a comparison of the data obtained from the perception input module with the digital track map data. the incident analysis module is designed to process the data related to the objects evaluated by the map matching module as potentially unwanted and additional information from detectors. the data are 8 https://doi.org/10.14311/app.2022.35.0008 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 35/2022 digital track map for the vexa expert system figure 1. logical architecture of the vexa system [1] analysed using a set of deterministic rules forming a nested multi-criteria expert system. on the basis of classification, it decides on the choice of reaction that is necessary to prevent or mitigate the impact of a potential incident. finally, the incident solver module is responsible for performing the correct reaction to possible incidents that the incident analysis module was unable to eliminate. the module includes relevant rules that define how to react to a given type of accident. as regards the vexa system surroundings, the perception subsystem includes various detectors providing data to the perception input module. among others, it is the object detector and the detectors of train control management system. the interface of other modules to various railway systems is provided by the vexa – train systems communication hub and vexa – trackside communication hub. the map matching module is also supposed to receive data from global navigation satellite system and to communicates with the digital track map. the logical architecture of the vexa system is shown in the figure 1. at the same time, the expected data flows are shown there [1]. 3. recognition of potentially dangerous objects as already outlined, the purpose of the digital track map is to provide the vexa system with the data describing the infrastructure used to operate the autonomous vehicle and its intermediate surroundings. even though the digital track map is not a part of vexa itself, it is being developed in close collaboration with the system in order to provide it with required inputs. whereas one of the vexa system goals is to detect unwanted objects around the railway line and especially on the running track, the map serves as a basis for comparison against the scene recognised on the basis of inputs captured by the means of onboard devices. while these inputs are supposed to be processed by the perception input module of the vexa system, the map matching module is responsible to compare the map with the recognised scene. for this reason, the map is supposed to include the location of expected objects around the railway line which are to be evaluated as harmless when being compared to the scene. the map matching module evaluates whether detected objects need to be classified as potentially dangerous. this case occurs if there is no corresponding data object included in the digital track map while a real object is detected by train sensors in a specific location. the information about the potentially dangerous object is to be provided the incident analysis module to decide on and carry out an appropriate reaction which affects the behaviour of the train. in the event of an irregularity being detected that cannot be solved by the incident analysis module and an accident threatens to occur, the incident solver module must also receive the appropriate data to take action [1]. 4. requirements arising from vehicle localization aspects considering the aspect of topological decomposition, the digital track map consists of two subsystems: digital map trackside and digital map onboard. the digital map onboard subsystem comprise the data tiles being loaded from the digital map trackside part depending on the specific train path. it works in close collaboration with the map matching module, responsible for further data processing. selection of individual data tiles to be loaded is based on the knowledge of the vehicle location [1]. the vehicle location information is also necessary to be known in order to determine the distance and direction in which the individual objects recorded in the digital track map are expected to be located 9 adam hlubuček acta polytechnica ctu proceedings relative to the vehicle. in order to calculate the values of these quantities, it is necessary to express the object location and the location of the vehicle in a common coordinate system. whereas linear coordinate systems are not suitable for this purpose (as they must always be assigned to a specific line or track) and there is a need to keep constant lengths, angles, and areas for measurement, it is necessary to express location through one of projected coordinate systems. it may require the recalculation of coordinates, in some cases. a projected coordinate system is defined on a flat, two-dimensional surface and is always based on a geographic coordinate system that is based on a sphere or spheroid. in a projected coordinate system, locations are identified by x and y coordinates. in addition, the z coordinate determining the elevation value can be expressed using a vertical coordinate system [3, 4]. as regard the railway vehicle, the location information is supposed to be provided by the gnss technology. it can be also determined using an odometry system of the vehicle, e. g. using collaboration with some of automatic train control or automatic train operation systems. the odometry information is especially important in such cases for which the location determined on the basis of gnss is not sufficiently accurate, in particular with regard to ensuring safetyrelevant functions. the location of a train cannot be determined by the means of the gnss technology in cases such as when the train is in a tunnel. gnss is also not able to reliably distinguish on which track in the station or on the multi-track line the vehicle is moving. this can only be determined in combination with route setting information provided by the railway traffic management system. for these reasons, a spatial description of the track guidance is also required to be included in the digital track map. this makes it possible to convert the odometric information related to the appropriate route into x, y and z coordinates. the same applies to descriptive data directly associated to individual points or segments defined within the network expressed at the level of individual tracks. 5. vexa and ato possible cooperation in the foreseeable future, vexa should be able to receive data from the avv system, which is the instance of ato used in the czech republic. the avv automatic train operation system was developed during the 1990s and put into revenue operation in 2000. it is a representative of the goa2 grade of automation system, which means the semi-automated train operation. the train is driven automatically, but the driver is still in the cab, e. g. to control the absence of obstacles in the track [2, 5, 6]. avv provides automatic driving for acceleration, cruise control and braking, comparing the actual speed with that requested by the driver and applying traction or braking to control the speed. in some cases, avv is able to work in close cooperation with the etcs automatic train control system. in such cases, these systems complement each other. avv can also use the same train parameters as etcs when calculating braking curves [6]. based on the principle of interconnection between the ato and the atc system, standardised european solution called ato over etcs (aoe) is being developed. this narrow harmonised cooperation brings also new concept of infrastructure data processing. instead of using fixed track map stored in avv, aoe uses more flexible concept. the description of the railway network corresponds to the decomposition into so-called segment profiles to which the relevant data relate. each running track is usually covered with several such segment profiles. the onboard ato part receives these data dynamically from the ato trackside subsystem. because the vexa system has the possibility to get some infrastructure data directly from ato, it is not needed to include them to the digital track map. the segment profiles, however, do not contain relevant spatial information which is necessary due to the need of recalculation of coordinates for the purposes of the digital track map and the map matching module functionalities. in order to localize the relevant data in geographic space, it is necessary to link individual segment profiles with appropriate geographically located elements. such a connection will make it possible to express the location of a point defined by the segment profile identifier and the depth of the segment profile penetration (absolute position from its beginning) also through geographical coordinates. this is essential both in terms of the ato infrastructure data localization and for the purpose of expressing the railway vehicle geographical location based on the odometry information [6]. 6. railtopomodel and gis principles application spatial expression of a railway transport route, which is the basis for localization of many infrastructure data description components, can take several different forms. the uic railtopomodel initiative aims to unify the various approaches to modeling railway infrastructure under one generic model allowing various user extensions managed in a common way. since the railtopomodel has become the irs 30100 solution in 2016, its application in various projects is gradually expanding. among others, the railml® 3 xml-based railway data format implements the concepts of the railtopomodel for data exchange purposes [7]. taking into account the railtopomodel principles, the multipurpose railway infrastructure model is being developed at the railway laboratory workplace at the czech technical university in prague 10 vol. 35/2022 digital track map for the vexa expert system figure 2. individual points of the track obtained through geodetic survey visualised in the arcgis pro software – faculty of transportation sciences. among several other projects, the digital track map represents a field for implementation and further development of this model. the multipurpose railway infrastructure model allows the user to view the recorded data in the context of individual objects or object classes, at the same time, however, it retains the form of a relational database, which allows it to be integrated into the geodatabase working under the gis software used [8]. as the platform for creating the digital track map, the arcgis pro software has been chosen. the software allows to open and edit various data formats from which the most convenient for storing data for the digital track map purposes is the file geodatabase format. it offers a comprehensive data repository integrating feature data sets, standalone tables and some other types of data sources. thanks to that, individual data records are not necessarily bound to specific gis features but they can simply be joined with them [4]. considering the principles of the railtopomodel recommendation, the topology of the digital track map network is designed using net elements and net relations. due to the nature of the task solved, the only possible approach is to use the micro level of detail with linear elements representing individual tracks interconnected by net relations at nodes representing switches. each linear element must be terminated in such a node and connected to the following linear elements with the use of a net relation [9, 10]. the digital track map, as it is created in the arcgis pro software, allows individual objects to be visualized and handled as standard gis features, i. e. points, lines and areas. therefore, the railtopomodel linear element class can be considered as a specialization of the line feature class, in the gis context. individual linear elements can therefore be represented and data-described as line features of the linear element feature class. these line features are fundamentally linestrings (sometimes also called polylines), i. e. sequences of points localized using geographical coordinates which can also have their altitude and stationing defined. in the case they represent linear elements, these points can be obtained through geodetic survey (as shown in the figure 2) or by sampling the analytical description of the track guidance (which is shown in the figure 3). when a geodatabase is taken into account, there is also the possibility to use the analytic description directly to define the line features (or parts thereof). arcgis pro offers the creation of arc and spiral segments in this way, even by entering their design parameters. it is necessary to consider whether such an option would be beneficial in the creation of the digital track map [4]. the net elements defined in this manner are supposed to become the basis for the location of various types of objects and properties of the railway infrastructure. among others, the aforementioned ato segments are assumed to be mapped to them in order to provide relevant data related to respective track sections. according to the railtopomodel terminology, railway infrastructure objects and properties which can be localized are collectively called net entities. after all, objects that are not directly related to the transport route can be considered as net entities, as well. this 11 adam hlubuček acta polytechnica ctu proceedings figure 3. horizontal curves rendered on the basis of the analytical description of the track geometry also make it possible to include among net entities those objects that are intended for comparison with the detected scene by the map matching module. net entities can be localized basically in two different ways. taking into account the class names defined in the multipurpose railway infrastructure model, it is possible to distinguish between physical and functional location. the physical location essentially corresponds to the localization with the use of the standard gis features and expresses the real location of physical objects in geographical space. such a location is therefore independent of the railway infrastructure and is not associated with any net elements. on the other hand, the functional location expresses the projection of the relevant object into associated point or segment defined within an (in our case linear) element or several (often consecutive) linear elements. this projection is typically perpendicular but in justified cases may be defined in another way to suit the functional nature of the localized object in relation to the railway transport route. depending on the method of functional location, we can further distinguish between spot, linear and area location. this division is based directly on the definition of the raitopomodel classes designed to describe the location aspects [9, 10]. as regards the digital track map, taking into account available data sources, the spot location is to be used for locating such entities as signals and balises, while the linear location is supposed to be used by horizontal curves, vertical curves, segment profiles, platform edges, level crossings and possibly other types of net entities. the area location could be used, for example, to locate turnouts, if it is decided that it is beneficial for the vexa project to record data about them. the main significance of the physical location from the point of view of the digital track map project is seen for the purposes of the railway surroundings objects which have no relation to a particular track. exclusively this type of location is supposed to be used, in such cases [1, 11]. 7. conclusions this article outlined possible approaches to creating the digital track map for the vexa project and introduced the basic issues which must be taken into account when defining the map feature classes. after the design of its structure is completed, the digital track map is supposed to be first used on the vexa system trial implementation on the čížkovice – obrnice railway line, also known as švestková dráha. the first testing is planned for the line section between the libčeves railway station and the židovice railway stop. the prospective intention is to extend the technology to a wider range of railway vehicles and parts of the railway network. acknowledgements the research presented in this paper is supported by the technology agency of the czech republic (ta čr) under project number ck01000111 "development of a train expert system for autonomous train behaviour". the author would like to thank ing. vítězslav lansfeld and ing. michal novák from ažd praha s. r. o. for valuable recommendations provided as part of the paper final review. references [1] m. matowicki et al. průběžná zpráva o činnostech, provedených v projektu vexa v roce 2020 [online]. tech. rep., čvut v praze – fakulta dopravní, 2021. [2] shift2rail. shift2rail multi-annual action plan, 2015. [3] arcmap. what are projected coordinate systems? [online], ©2016. https://desktop.arcgis.com/en/ arcmap/10.3/guide-books/map-projections/ about-projected-coordinate-systems.htm. [4] t. corbin. arcgis pro 2.x coockbook. packt publishing, 2018. [5] a. lieskovský, i. myslivec. etcs a ato – poprvé společně. reportér ažd praha (3): p. 46 – 48, 2011. [6] a. lieskovský, i. myslivec. ato and etcs can work together. railway gazette international p. 19 – 22, 2011. 12 https://desktop.arcgis.com/en/arcmap/10.3/guide-books/map-projections/about-projected-coordinate-systems.htm https://desktop.arcgis.com/en/arcmap/10.3/guide-books/map-projections/about-projected-coordinate-systems.htm https://desktop.arcgis.com/en/arcmap/10.3/guide-books/map-projections/about-projected-coordinate-systems.htm vol. 35/2022 digital track map for the vexa expert system [7] railtopomodel. railtopomodel [online], ©2021. http://www.railtopomodel.org. [8] m. leso, d. kamenický, p. koutecký, a. hlubuček. dopravní sál fd – železniční laboratoř pro výuku i výzkum. vts správy železnic 1, 2019. [9] uic. railtopomodel v1.0 [pdf], 2016. http://www.railtopomodel.org/en/download/ irs30100-apr16-7594bca1524e14224d0.html?file= files/download/railtopomodel/180416_uic_ irs30100.pdf. [10] uic. railtopomodel v1.1 [online], 2017. https://uic.org/rail-system/railtopomodel. [11] ažd. zdd [online], 2019. https://www.azd.cz/cs/ dalsi-informace/provoz-drahy/zdd. 13 http://www.railtopomodel.org http://www.railtopomodel.org/en/download/irs30100-apr16-7594bca1524e14224d0.html?file=files/download/railtopomodel/180416_uic_irs30100.pdf http://www.railtopomodel.org/en/download/irs30100-apr16-7594bca1524e14224d0.html?file=files/download/railtopomodel/180416_uic_irs30100.pdf http://www.railtopomodel.org/en/download/irs30100-apr16-7594bca1524e14224d0.html?file=files/download/railtopomodel/180416_uic_irs30100.pdf http://www.railtopomodel.org/en/download/irs30100-apr16-7594bca1524e14224d0.html?file=files/download/railtopomodel/180416_uic_irs30100.pdf https://uic.org/rail-system/railtopomodel https://www.azd.cz/cs/dalsi-informace/provoz-drahy/zdd https://www.azd.cz/cs/dalsi-informace/provoz-drahy/zdd acta polytechnica ctu proceedings 35:8–13, 2022 1 introduction 2 logical architecture of the vexa system 3 recognition of potentially dangerous objects 4 requirements arising from vehicle localization aspects 5 vexa and ato possible cooperation 6 railtopomodel and gis principles application 7 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0255 acta polytechnica ctu proceedings 38:255–261, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague improving the properties of unburned earth to control the interior microclimate michal procházka∗, zbyněk zušťák, pavel svoboda czech technical university in prague, faculty of civil engineering, department of construction technology, thákurova 7/2077, 166 29 prague, czech republic ∗ corresponding author: prochmi2@fsv.cvut.cz abstract. the article concerns the problematics of unburned earth used as material for moulding formwork or clay plaster with lowered contraction, especially about interior unburned earth plaster with the heightened percentage of clay and therefore improved sorption abilities. by adding a plasticizer, it is possible to modify the fluidity of the mixture. with lessening the amount of water it is possible to lower the shrinkage. furthermore, we can increase the ratio of clay to sand and thus we can control the indoor microclimate. this is because the clay provides adsorption properties. unburnt clay has been widely used as a building material in the past around the world. in the last century, this material was rather despised. nowadays, however, this trend is reversing and the unburnt clay is experiencing a renaissance. increasingly people emphasize ecology, a healthy and balanced indoor climate, and demand buildings with low operational and energy intensity. all this can be offered by unburnt clay, because it is energy-efficient in production, easily accessible in almost all parts of the world, harmless, and beneficial to human health, as clay can contribute to the improvement of the internal microclimate to a much greater extent than other building materials, as confirmed by many scientific findings. keywords: earth, clay, microclimate. 1. introduction unburned earth can be defined as building material held together primary by clay. unburned earth is a common material in the building industry around the world. but the european building norms don’t know this material. unburned earth can be defined as soil or a mixture of soils with added ingredients suitable for building purposes which compose of fine-grained fractions of clay particles as the primary binder and dust particles, further composed of sand and gravel fraction used as filler. also other materials can be used in the mixture such as scattered reinforcement formed by a fibrous admixture, pigment and secondary binders and stabilizers, all used in the right amount. the primary source of the binding ability is clay materials. clay is also the main source of the positive features as sorption, especially the regulation of the indoor microclimate. depending on the physicochemical properties, it occurs mainly in the kaolinite group, the montmorillonite group and the illite group. these clay minerals are often forming mixed clay sediments. the structure of clay minerals is layered. it consists of layers of oxygen tetrahedra with a silicon atom in the middle and layers of oxygen octahedra with an aluminum atom in the middle and an interlayer space (see figure 1). water penetrates into this interlayer space, especially in the case of montmorillonibte, which causes swelling. swelling and thus shrinkage during drying is the lowest in kaolinite, higher in illite and highest in montmorillonite. if the water recedes from this intermediate layer, the electrostatic attractive forces, which are the essence of the binding ability of clay minerals, will increase. it is, therefore, not a chemical binder and the process is reversible. water is also important when moving the components of unburned earth, including mineral plates and their arrangement during drying. here the water can be partially replaced. kneading the mixture or vibration helps to ensure a good arrangement. this can reduce the required amount of water needed for the mixture. however, these procedures are mostly associated with the technology used, and they can often not be replaced or used at all, especially for plasters. in terms of soil classification according to čsn en iso 14688-2 [1], the result is most often soils sisa, clsa, grsisa, grclsa, grsasis, grsacls, sagrsis and sagrcls. for practical reasons, they are often mixed from cl, sicl, clsi soils in practice called clay or loess and sa, grsa soils in practice called mortar sand. for some techniques, gr, sagr in practice called gravel or gravel sand is added. we can thus determine the ratio of individual components. by choosing clay minerals and choosing the ratio of individual components, we influence the binding and shrinkage during drying. in terms of soil granularity according to čsn en iso 17892-4 [2], clay represents a fine-grained fraction up to 0.063 mm consisting of clay and dust particles, mortar sand represents a sand fraction from 0.063 to 2 mm consisting of fine, medium and coarse sand particles and gravel fraction. from 2 to 63 mm consisting of fine, medium and coarse gravel particles. granulometry in the fine-grained fraction is not interesting 255 https://doi.org/10.14311/app.2022.38.0255 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en m. procházka, z. zušťák, p. svoboda acta polytechnica ctu proceedings figure 1. structure of clay materials. because it does not say whether and what type of clay particle it is. granulometry together with the shape of the particles is very interesting for the sand and gravel fraction of the filler. its suitable curve can positively affect drying shrinkage and strength. sorption is a general term for absorption, adsorption. absorption is the sorption of one substance into another in its entire volume. adsorption is the binding of a gaseous or liquid substance to the surface of another substance. there are two types of adsorption. physical adsorption arises from van der waals attractive forces. chemisorption is stronger than physical adsorption, it is formed by chemical bonds. in the context of unfired earth or clay particles, we speak of physical adsorption. 1.1. ratios of primary components wo2005092816 (a2) entitled “use of polymer powder compositions which can be redispersed in water for loam construction materials” discloses a clay dose of 10 to 30 % by weight [3]. bricks, blocks, mortars for masonry, plaster or are poured or smoked into formwork are made of unburned earth. the amount of clay typically depends on the type and origin of clay or earth, the content of clay fraction or clay minerals, the type of clay minerals, their binding ability and their effect on shrinkage or swelling in contact with water, granulometry of other particles, but especially on the purpose the use of a clay mixture and the associated dose of mixing water. in the case of the use of unburned earth as clay plasters, the approximate proportion of clay is from 10 to 15 % by weight. the proportion from 15 to 30 % by weight is suitable for clay mixtures with a lower water content for stamping into formwork. the values of the proportion approximately above 30 % are suitable at most for mixtures for the production of bricks, where shrinkage does not carry out such a role, because shrinkage takes place outside the structure. the use of a suitable fibrous admixture such as cuttings, shives, etc. makes it possible to positively increase the doses of clay. the clay dose is determined according to the tests below. the basic binder in unburned earth is clay. pure clay has a high binding capacity. the minimum binding capacity of the mixture is stated by vlastimil havlíček and karel souček in the book: “buildings made of unfired clay” as 250 g/5 cm2 [4]. binding capacity is tested on 8-shape samples with defined moisture (see figure 2). however, the clay shrinks during drying due to the evaporation of water. this can cause cracks in the unburned earth during drying. therefore, it is necessary to add dust and sand or gravel particles to the clay as a filler which reduces the total shrinkage to a tolerable limit of approximately 2 % in length. at higher total shrinkage, for example, cracks begin to form in clay plasters, and separation may also occur from less adhesive substrates. when measuring shrinkage, havlíček and souček [4] start from a poor initial consistency. this corresponds broadly to the consistency for steamed clay. in accordance with the procedures from the book “building with earth” by minke [5], the measurement of shrinkage must be based on the initial consistency for processing according to the technique, and thus on the amount of water in the mixture, which is always given technologically. for clay plasters, gernot minke states a consistency with a pour of 180 mm according to din 1060-3, which it further reduces to 140 mm [5]. in this article, total shrinkage means shrinkage, which is based on determining the consistency of fresh mortar according to čsn en 1015-3 [6] and the prescribed pour value according to čsn en 1015-2 [7] and thus for mixtures over 1200 kg/m3 pour value 175 ± 10 mm. spill values are measured on a shaker table (see figure 3). this value corresponds to measurements performed in practice, where several observations and samplings from several craftsmen were made just before the plaster was applied. the measurement of total shrinkage is performed either according to havlíček and souček [4] in steel or wooden forms measuring 25 × 40 × 220 mm on marks 200 mm distant or according to minke [5] in metal forms of the same dimensions according to din 18952 on marks 200 mm distant or in metal forms size 40 × 40 × 160 mm provided in the faces with pins for measurement according to čsn en 12617-4 [8] or by measuring the resulting length dimension of formed beams from metal forms size 40 × 40 × 160 mm 256 vol. 38/2022 improving the properties of unburned earth . . . figure 2. test of binding capacity figure 3. manual flow table. 257 m. procházka, z. zušťák, p. svoboda acta polytechnica ctu proceedings figure 4. test of shrinkage. according to čsn en 1015-11 [9] (see figure 4). the total shrinkage is measured from the above initial consistency of 175 ± 10 mm corresponding to the processing state to the stabilization of the shrinkage at the equilibrium humidity state of the unfired clay. the shape of the form does not affect the result. total shrinkage is expressed in %. the ratio between clay and other particles is thus more or less dependent on the technologically given amount of water in the mixture. another is the dose of water and therefore the consistency for bricks and steamed clay, and it is different for screeds and plasters that need more water for processing and shrinkage causes more problems. the result is that the smallest ratio of clay to other particles is put into the plaster. although this reduces the overall shrinkage, which is essential especially for plasters, it also reduces the binding ability or bonding and reduces other positive effects of the clay. so it is always a compromise between shrinkage and bonding. this compromise can be improved by suitable granulometry. clay in plasters is also a carrier of positive properties such as the adsorption of odours, dusts, allergens, but especially air humidity. moisture is adsorbed by the clay in clay plasters when there is an excess of ambient air and when it is deficient in the ambient air it is evaporated again. thus, clay plasters regulate air humidity in rooms. it is therefore desirable to produce plasters with a higher clay content, which is currently not possible due to shrinkage and process ability. 1.2. ingredients and admixtures further improvements can be achieved by replacing part of the filler with a fibrous admixture. this helps in addition to shrinkage reduction and a better and more even distribution of cracks. at the same time, it increases compressive and bending strengths which is especially desirable for plasters. fibres up to approx. 4 cm are suitable for plasters. higher doses of fibrous admixtures can impair processability. longer fibres and larger quantities can be used especially for lightweight structural or filling clay or for levelling layers. dosing is usually performed by volume. in laboratory conditions dosing is often complicated by the size of the forms. other additives such as secondary or stabilizing binders can be added to unburned earth, which can reduce the reaction to the effects of water or improve the strength characteristics or increase the adhesion to the substrate, as stated in cz document pv 2008663 entitled “unburned earth stabilized with polymers” [10]. bc. kateřina nováková mentions the use of plasticizers in her diploma thesis “research binding capacity of unburnt earth in time” [11] or fluxes that reduce internal friction. however, specific plasticizers are not mentioned, and reference is made in the literature below. the issue of processability of clay as a raw material for the production of fired ceramic goods is well known in the ceramic industry. certain analogies can be found between the ceramic and clay industries, but the nomenclature, composition of the clay and use differ. as gernot minke states on page 39: “in the ceramic industry, fluid thinning mediums are used to attain higher liquidity, thereby allowing less water to be used (in order to reduce shrinkage). typical thinning media are sodium waterglass (na2o · nsio2), soda (na2co3), and humus acid and tannic acid. tests conducted at the brl at the university of kasselshowed that these methods were of very little relevance to earth as a building material. but tests with whey were successful.” [5]. in the field of clay, cheap natural substances are traditionally used to reduce shrinkage, as stated by ivana žabičková in the book “hliněné stavby” on page 31: “je známo též užití koňské moči, která účinně eliminuje trhlinky a zvyšuje odolnost proti erozi.” [12] which means in english: it is also known to use horse urine, which effectively eliminates cracks and increases resistance to erosion. in the book “hliněné stavby novej generacie” [13] on page 83, petr suske mentions the use of recesses to reduce internal friction. 258 vol. 38/2022 improving the properties of unburned earth . . . 1.3. chemical additives for building materials in the concrete industry plasticizers are used to reduce internal friction. the highest effect is achieved with polycarboxylate-based plasticizers. tests of their use for unfired clay have shown their unsuitability. therefore, it is assumed that unfired clay cannot be combined with concrete plasticizers. 2. solutions the above-mentioned disadvantages of unburnt earth are solved by the unburnt earth with controlled shrinkage according to the invention “unfired clay with controlled shrinkage.”. according to it, the mixture contains 95 to 99.99 % by weight. unburnt earth and 0.01 to 5 wt. at least one plasticizer selected from the group: lignosulphonate-based, naphthalene-based and melamine-based [14]. 2.1. unburned earth for casting in the first step, using the addition of a plasticizer, we can increase the plasticity or flow ability of the unfired clay without increasing the shrinkage during drying. this can be used wherever the shape of unfired clay adds formwork. the process for producing unburned controlled shrinkage earth for casting is as follows: first, a dry mixture of unburned, shrinkage-controlled earth containing from 10 % to 30 % by weight is produced. clay by determining the proportion of clay in the unfired clay as a maximum in the range of 10 to 30 % by weight in the first step of the total shrinkage test, while maintaining the total shrinkage of the unburned earth up to 2 %, then in the second step by the fresh mortar consistency test using of the shaking table determines the optimal dose of plasticizer in the range of 0.01 to 5 % by weight, at which the spillage of unburned earth is maximum and in the third step the dose of plasticizer determined in the second step is made up to 100 % by weight. unburned earth with a clay content determined in the first step and into the resulting mixture in an amount ranging from 66.7 to 90 % by weight. water is mixed in an amount in the range of 10 to 33.3 % by weight, determined by a consistency test using a shaker table, so that the amount of water added to the original mixture from the first step corresponds to a pour of 175 mm [14]. 2.2. unburned earth for plasters with reduced shrinkage in the next step, in addition to the addition of the above-mentioned dose of plasticizer, we can reduce the amount of mixing water so that the consistency again corresponds to the plaster. this will create a plaster with reduced shrinkage during drying. the process for the production of unfired controlled shrinkage clay for plasters with reduced shrinkage is as follows: first, a dry mixture of unburned earth with controlled shrinkage containing from 7.5 % to 20 % by weight is produced. clay by determining the proportion of clay in the unburned earth as a maximum in the range of 7.5 to 20 % by weight in the first step of the total shrinkage test, while maintaining the total shrinkage of the unfired clay up to 2 %, then in the second step by the fresh mortar consistency test using a shaking table, the optimum dose of plasticizer is determined in the range of 0.01 to 5 % by weight at which the spillage of unburned earth is maximal and in the third step the dose of plasticizer determined in the second step is made up to 100 % by weight. unburned earth with the proportion of clay determined in the first step and in the resulting mixture in an amount ranging from 66.7 to 90 % by weight. admixed water in an amount in the range from 10 to 33.3 % by weight so that the shake table consistency test shows a spillage of 175 mm [14]. 2.3. unburned earth for plasters with increased sorption in the last step, the addition of the above-mentioned dose of plasticizer and the reduction of mixing water, which would otherwise cause a reduction in shrinkage, can be used to increase the ratio of clay to sand so that shrinkage is maintained. this will cause a plaster with increased sorption. clay is the main carrier of sorption abilities. the plaster thus helps to clean the air and especially to regulate the humidity microclimate. the process for the production of unfired controlled shrinkage clay for plasters with increased sorption is as follows: first, a dry mixture of unburned, shrinkagecontrolled earth containing from 20 % to 50 % by weight is produced. clay by determining the proportion of clay in the unburned earth as a maximum in the range of 19.99 to 49.99 % by weight in the first step of the total shrinkage test, while maintaining the total shrinkage of the unburned earth up to 2 %, then in the second step by the consistency test fresh mortar using a shaker table determines the optimal dose of plasticizer in the range of 0.01 to 5 % by weight, at which the spillage of unfired clay is maximum, then in the third step the total shrinkage test is performed on the mixture of unburned earth according to the first step and plasticizer according to the second step, determines the increased proportion of clay as a maximum in the range of 20 to 50 % by weight, while maintaining the total shrinkage of this mixture up to 2 %, and in the fourth step the dose of plasticizer determined in the second step is made up to 100 % by weight. unburned earth with an increased proportion of clay, determined in the third step and into the resulting mixture in an amount ranging from 66.7 to 90 % by weight admixes water in an amount in the range from 10 to 33.3 % by weight so that the shake table consistency test shows a spillage of 175 mm [14]. 2.4. results of experiments figure 5 shows an approximate behavior curve of a melamine-based plasticizer in a clay mixture. 259 m. procházka, z. zušťák, p. svoboda acta polytechnica ctu proceedings figure 5. evaluation of the effect of melamine-based plasticizer [15]. figure 6. shrinkage of clay mixtures [15]. the k mixture is a control mixture, with excessive shrinkage and therefore with an excessive dose of clay. commonly we would reduce the ratio of clay to sand. but this is not desirable from the point of view of the microclimate. the pl00 mixture has a lower water dose and a melamine-based plasticizer is metered into the other pl mixtures. the reaction of the clay plaster mixture to the gradual addition of a melamine-based plasticizer can be observed on the blue curve. figure 5 shows that the plasticizer had the highest efficacy at a dosage of 0.2 %. after optimizing the amount of mixing water, with a final spill of 184 mm, we can observe a decrease of 15 % [15]. in figure 6, the mixtures are assigned the appropriate shrinkage in percent. the maximum acceptable shrinkage for earthen plasters is 2 %. the results show the possibilities to increase the fluidity of the mixture, further reduce water by 15 % and reduce shrinkage from 4.38 % to 1.88 %. this can be used to increase the clay to sand ratio and thus control the interior microclimate. 3. conclusion unfired clay is widely used in ecological construction. it is suitable for the preparation of unburned earth for casting and plaster. it is then used in the production of external and internal plasters with reduced shrinkage during drying. it will then use the unique properties in the production of internal plaster with increased sorption. this makes it possible to better regulate the cleanliness and humidity of the indoor air and thus have a beneficial effect on the microclimate. the specific resulting effect on the interior microclimate will be the subject of further research. however, it is already certain that the ratio of clay to sand can be increased without increasing shrinkage, which is limiting for clay plasters. acknowledgements we would like to thank the co-author of the patent pavlína znamenáčková, who contributed to the creation of the unfired clay with controlled shrinkage patent with her laboratory work. 260 vol. 38/2022 improving the properties of unburned earth . . . references [1] čsn en iso 14688-2, geotechnický průzkum a zkoušení – pojmenování a zatřiďování zemin – part 2: zásady pro zatřiďování, únmz, praha, 2005. [2] čsn en iso 17892-4, geotechnický průzkum a zkoušení – laboratorní zkoušky zemin part 4: stanovení zrnitosti, únmz, praha, 2005. [3] wacker polymer systems gmbh & co. kg. use of polymer powder compositions that can be redispersed in water for loam construction materials. patent. inventors: h. lutz, a. sommerauer, k. bonin. priorities: 2004-03-18. mpt: c04b28/00, c04b24/26, c04b24/38. application: de200410013468. published as: wo2005092816. german patent and trade mark office. [4] v. havlíček, k. souček. stavby z nepálené hlíny. státní zemědělské nakladatelství, praha, 1958. 98 p. [5] g. minke. building with earth. birkhäuser, berlin, 2006. 208 p. isbn 978-3-7643-7873-8. [6] čsn en 1015-3, zkušební metody malt pro zdivo – part 3: stanovení konzistence čerstvé malty (s použitím střásacího stolku), únmz, praha, 2000. [7] čsn en 1015-2, zkušební metody malt pro zdivo – part 2: odběr základních vzorků malt a příprava zkušebních malt, únmz, praha, 1999. [8] čsn en 12617-4, výrobky a systémy pro ochranu a opravy betonových konstrukcí – zkušební metody – part 4: stanovení smršťování a rozpínání, únmz, praha, 2003. [9] čsn en 1015-11, zkušební metody malt pro zdivo – part 11: stanovení pevnosti zatvrdlých malt v tahu za ohybu a v tlaku, únmz, praha, 2000. [10] m. procházka, p. svoboda. unfired clay stabilized with polymers. patent. inventor: m. procházka. priorities: 2008-10-23. mpt: b28c1/04, c04b14/10, c04b24/24. application: cz2008663a. published as: cz302139b6. industrial property office. [11] k. nováková. research binding capacity of unburnt clay in time. master’s thesis, czech technical university in prague, praha, 2017. [12] i. žabičková. hliněné stavby. era, brno, 2002. 174 p. isbn 8086517217, 9788086517216. [13] p. suske. hlinené domy novej generácie. alfa, bratislava, 1991. 159 p. isbn 8005008945, 978-8005008948. [14] czech technical university in prague. unfired clay with controlled shrinkage. patent. inventors: m. procházka, p. znamenáčková. priorities: 2018-03-22. mpt: c04b14/10, c04b24/18, c04b24/12. application: cz2018143a. published as: cz307890b6. industrial property office. [15] p. znamenáčková. využití plastifikátorů při výrobě hliněných omítek. master’s thesis, czech technical university in prague, praha, 2017. 261 acta polytechnica ctu proceedings 38:255–261, 2022 1 introduction 1.1 ratios of primary components 1.2 ingredients and admixtures 1.3 chemical additives for building materials 2 solutions 2.1 unburned earth for casting 2.2 unburned earth for plasters with reduced shrinkage 2.3 unburned earth for plasters with increased sorption 2.4 results of experiments 3 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0429 acta polytechnica ctu proceedings 38:429–435, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague development of a realistic scenario for the thermal energy demand of residential buildings in bavaria till 2050 michael barton∗, franziska pichlmeier, marissa wolter, christian schweigler munich university of applied sciences, cenergie – center for energy-efficient buildings and districts, lothstr. 34, munich, 80335, germany ∗ corresponding author: michael.barton@hm.edu abstract. to achieve the climate protection goals, enormous efforts must be undertaken in all sectors: households, industry, commerce, and transport. following a straightforward approach, the political objectives foresee, that in the building sector a massive reduction of energy use for the heat supply is accomplished. however, previous investigations have shown that this reduction of energy consumption is not feasible. older buildings exhibit especially high energy demand and emissions. yet, due to the low refurbishment rate, no substantial change of the heat demand can be expected within the next decades. by fully renovating the entire residential building stock, approximately 70 % of the final energy demand and related co2 emissions could be saved, still not enough to reach the political goals. therefore, renovation or renewal of buildings and of the use of renewable energy sources have to be implemented jointly for achieving the desired savings. the methods used to estimate the characteristics of bavaria’s residential building stock as well as its heating energy demand and related co2 emissions for the year 2050 are presented. alternative goals are given which base upon the achievable final energy saving for a realistic renovation scenario accompanied by further reduction of co2 emissions by using renewable energies. keywords: refurbishment, building stock, climate neutrality. 1. introduction in order to achieve the european, german and bavarian climate protection goals, enormous efforts must be undertaken to accomplish substantial emissions reduction in all sectors: households, industry, commerce, and transport. following a straightforward approach, the political objectives foresee, that in the building sector this goal is reached by massive reduction of the final energy use for the heat supply of buildings. however, previous investigations have shown that the targeted reduction of final energy consumption in the building sector is not feasible. older buildings exhibit especially high energy demand and high co2 emissions. yet, due to the low refurbishment rate, no substantial change of the heat demand is to be expected in the near future. by fully renovating the entire residential building stock, approximately 70 % of the final energy demand and related co2 emissions could be saved, still not enough to reach the political goals. therefore, goals that achieve reduction of final energy and co2 emissions through a combination of renovation or renewal of buildings and the use of renewable energy sources have to be defined in order to meet reasonable objectives regarding co2 emissions. the proposed pathway could form the starting point for reformulating the political goals for the future energy use, finally targeting climate neutrality of the building heat supply sector. 2. methods to define realistic objectives for an effective transformation of the energy supply to residential houses, a detailed characterization of the heat demand with regard to all relevant criteria is required. there are numerous sources that describe building age classes as well as building types and provide statistical data on flat size, flat layout, specific energy demand and other parameters. however, there is no comprehensive overview available of how the individual criteria influence the heat demand of a specific building. therefore, data regarding the composition of the bavarian residential building stock and its final energy demand has been collected and the impact of different criteria has been shown. as final result a very detailed data set comprising 840 building variants, classified by four criteria, has been obtained from which a large number of specific statements can be derived. as shown in the figures 1 to 4, these data were generated by linking and correlational analysis of individual data sources in order to derive the desired statistical results for the characterization of the building stock and its energy demand for space heating and domestic hot water. at the beginning, the demand of final energy for the heat supply of the bavarian residential building stock has been analysed. in a first step, the perspective of the energy supply has been taken to obtain a reference value for the heat supply data, to be derived from the 429 https://doi.org/10.14311/app.2022.38.0429 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en m. barton, f. pichlmeier, m. wolter, c. schweigler acta polytechnica ctu proceedings figure 1. balance heat supply residential buildings in bavaria 2016. ([1–3]). physical data of the building stock in the following step. (figure 1). for this purpose, it was determined which amounts of the different energy carriers are provided for the heat supply of the residential houses. in the second step, the situation has been investigated from the perspective of the consumers (figure 2). the composition of the bavarian residential building stock has been characterized first, then the final energy demand has been determined, partitioned according to various criteria such as building categories and age. the main part of the physical data comes from the german census [4, 5] and the european wide developed building typology within the tabula project [10, 11]. the energy values of the two approaches – supply and consumption – are congruent with a deviation of less than 5 %. subsequently, a forecast for the development of the bavarian residential building stock and its final energy demand for the period until 2050 has been generated (figure 3). for this purpose, assumptions have been made in a realistic scenario about how the construction of new residential houses, the demolition and the renovation rate for existing buildings and the specific energy demand of the different building types will develop in the future. results for the energy demand and greenhouse gas emissions have been generated with reference to the climate policy objectives of the eu, the federal republic of germany and the federal state of bavaria. these reduction goals were then transferred to the heat supply of the bavarian residential building stock (figure 4), comprising the heat quantities for space heating and domestic hot water. in order to evaluate the reduction potential, it was considered how far the final energy demand of the residential building stock could be reduced until 2050, if all existing buildings were fully refurbished and the new houses had the highest energy standard of the respective year of construction. this shows the theoretical maximum savings potential in final energy and the associated greenhouse gas emissions for the heat supply of the bavarian residential building stock. 3. results and discussion detailed results and analysis on the composition of the bavarian building stock and its energy demand for space heating and domestic hot water preparation have already been published [25, 26]. examples for the use of these data for further analysis supported by appropriate reference building models have been presented in [27, 28]. a rough summary of the results for the heat demand concerning the shares of different building types are shown in table 1. figure 5 and table 2 show the results on the final energy values for different scenarios. political objectives: transferring the global political objectives for the energy use in all sectors to the heat demand (space heating and domestic hot water) of residential houses results in a target value for the final energy of approx. 24 800 gwh. compared to all other scenarios, this represents the politically targeted energy demand of the building sector in 2050. rate of refurbishment constant: a constant renovation rate of 1 % per year was assumed for this “business as usual” case. in this scenario, the energy demand is reduced by 19 % from about 101 300 gwh in 2016 to about 82 000 gwh by 2050. although it is not a pessimistic scenario, the result for the remaining energy demand is more than 57 000 gwh/a above the political goal. rate of refurbishment enhanced: this scenario based on optimistic, yet still realistic assumptions, for the development until 2050 still misses the political objectives significantly. the reduction of the final energy demand by approx. 30 400 gwh to approx. 71 000 gwh is not sufficient. an additional reduction of more than 46 200 gwh to about 24 800 gwh would be necessary to reach the political goal. to reach this goal the effort for renewal and renovation of the building stock would have to be increased by a factor of about 2.5. full refurbishment: the maximum potential for the reduction of the final energy demand can be exploited through the full refurbishment of the residential build430 vol. 38/2022 development of a realistic scenario for the thermal energy . . . figure 2. balance heat demand residential buildings in bavaria 2016. ([4], [5], [6–8], [9–11], [12, 13]). figure 3. balance heat demand residential buildings in bavaria 2050. ([6–10, 14, 15], [9–11]). figure 4. objectives for the heat supply of residential buildings in bavaria until 2050. ([14, 16–20], [2, 21, 22], [23, 24]). 431 m. barton, f. pichlmeier, m. wolter, c. schweigler acta polytechnica ctu proceedings residential building stock bavaria 2050 buildings final energy [pcs] [gwh] residential buildings total 3 501 476 100 % 70 895 100 % buildings unrefurbished 1 307 604 37.3 % 36 706 51.8 % buildings partly refurbished 996 197 28.5 % 23 591 33.3 % buildings fully refurbished 1 197 675 34.2 % 10 598 14.9 % detached houses 2 347 775 67.1 % 46 308 65.3 % semi-detached houses 495 207 14.1 % 7 915 11.2 % terraced houses 515 769 14.7 % 13 246 18.7 % other building types 142 725 4.1 % 3 427 4.8 % buildings with 1 flat 2 358 452 67.4 % 33 100 46.7 % buildings with 2 flats 624 305 17.8 % 14 251 20.1 % buildings with 3–6 flats 334 530 9.6 % 10 352 14.6 % buildings with 7–12 flats 128 601 3.7 % 7 611 10.7 % buildings with 13 or more 55 588 1.6 % 5 582 7.9 % table 1. rough summary of the results for the composition of the bavarian building stock and its energy demand for heating supply. the data for 2050 shows the scenario “rate of refurbishment enhanced”. figure 5. demand and objective values for final energy and for co2 emissions. scenarios 2050 renewable share renewable expansion final energy total final energy final energy resulting co2 reduction based on a heat demand [gwh] [gwh] [gwh] [million t co2] of 101 315 gwh in 2016 (based on 2016) (based on 2050) (based on 2016) rate of refurbishment constant 82 083 (81 %) 64 143 (78 %) +37 396 +140 % 8.27 rate of refurbishment forced 70 895 (70 %) 52 955 (75 %) +26 208 +98 % 5.80 full refurbishment 30 330 (30 %) 7 12 390 (41 %) -14 357 -54 % -3.17 political objective final energy 24 802 (24 %) 6 862 (28 %) -19 885 -74 % -4.40 table 2. requirements to achieve co2 objectives of 3.9 million t co2/year. 432 vol. 38/2022 development of a realistic scenario for the thermal energy . . . ing stock. yet, it is practically impossible to achieve this level of refurbishment. among other aspects, there is not sufficient capacity available, such as materials and workforce, to realize this scenario. and for a variety of reasons, some buildings cannot reach the highest insulation standard. the protection of historic buildings can be given as an example. although not realistic, this scenario provides an answer to the question whether the political objectives are theoretically achievable by refurbishment measures. in this case, the final energy demand of the building stock in 2050 is reduced by approx. 70 % to approx. 30 300 gwh, compared to the situation in 2016. thus, this scenario comes closest to the political objective of 24 800 gwh final energy consumption. the analysis leads to the conclusion that the political objectives for the final energy demand of the bavarian residential building stock cannot be achieved by refurbishment alone, taking into account the current state of building renovation and heating technology. yet, with regard to the current composition of the energy supply in the bavarian building sector, as shown in figure 5 (left section: 2016), it becomes obvious that the political goals have to be re-oriented. currently, approx. 26 750 gwh of the final energy demand for heating is covered by renewable energy sources. this amount exceeds the remaining energy demand, which is targeted by the political goals, exclusively relying on energy saving measures. thus, it has to be concluded that the future pathway for the energetic development of the building sector has to focus rather on the emission of the greenhouse gas co2 than on the building energy demand only. and in return, development scenarios have to incorporate both improvement of the thermal quality of the building structure and variation of the use of fossil and renewable energy sources. for this perspective, a target value for 2050 for the reduction of the co2 emissions related to the heat supply in the bavarian building sector has been calculated, independently from the limit found for the energy demand, as described in figure 4. by applying the politically targeted co2 reduction rate to the bavarian residential building sector, the emissions related to the heat supply have to be limited to 3.97 million tons per year. in order to assess the required contribution of renewable energy for the different refurbishment scenarios described above, it has been assumed that in 2050 fuel oil is fully banned and only natural gas will be used as a fossil energy carrier. the results of this investigation are shown in table 2 and figure 5. for the “business as usual” scenario with constant refurbishment rate, it is found that the use of renewable energy sources has to be increased by 37 400 gwh, equivalent to an increase by 140 % relative to the current value of 26 750 gwh. additional co2 savings of 8.3 million tons would be accomplished in order to reach the co2 emission goal, calculated with the actual emission factor for natural gas. in total, the renewable share would cover 78 % of the final energy demand of the building heat sector in 2050. in case of the “enhanced refurbishment scenario” the contribution of renewables has to be increased by 98 % in comparison to the current status (2016), covering 75 % of the total demand. as stated above, in the case of full refurbishment the objective for final energy would be missed. yet, a rather low absolute amount of renewable energy (12 390 gwh), equivalent to 41 % of the remaining final energy demand, is required to fulfil the objective for co2 emissions. apart from the dramatic reduction of the building energy demand, this scenario could represent a viable transformation path of the energy supply, since the required change of the building stock will in return strongly influence the potential volume for utilization of renewable energy sources. although the absolute amount of renewable energy would decrease by about 54 % compared to 2016, the renewable share would increase from 26 % in 2016 to 41 %. if both objectives shall be fulfilled – final energy, and co2 emissions – 28 % of the remaining demand of final energy are to be covered by renewable sources. from these considerations, it can be concluded that a singular focus on the final energy goal does not provide a viable solution, while concentrating on the objective value for co2 emissions is expedient. as shown in figure 5, by combining efforts in refurbishment of the building stock plus enhanced use of renewable energy sources the goal for co2 emissions can be achieved. to identify a realistic pathway, future development should achieve the maximum possible saving of final energy and accomplish the remaining reduction in greenhouse gas emissions by expanding the use of renewables. this strategy is best represented by the “enhanced rate of refurbishment” scenario. 4. conclusions a method to generate detailed statistical data of the bavarian building stock, its energy demand for heat supply (space heating and domestic hot water) and their future development until 2050 with regard to the political objectives of the eu, the federal republic of germany and the federal state of bavaria have been described. the procedure for creating a realistic scenario to reach the goals until 2050 has been explained. the theoretical maximum potential for saving final energy and reducing the associated greenhouse gas emissions for the heat supply in the bavarian building sector through full refurbishment and construction of new buildings with the highest energy standard until 2050 has been shown. based on the analysis of the possible development of the building structure, it has been concluded that political objective solely related to the building energy demand are unachievable. as an alternative, technically feasible goals have been worked out with focus on reducing co2 emissions as extensively as possible 433 m. barton, f. pichlmeier, m. wolter, c. schweigler acta polytechnica ctu proceedings by renovating residential houses and achieving further savings through the increased use of renewables. a realistic pathway has been presented, comprising a reduction of the final energy demand of the bavarian building sector by 30 % accompanied by a 75 % coverage of the remaining demand by renewable sources. as central result, the co2 emission of the building stock in 2050 is limited to 3.97 million tons per year, representing a reduction by 78 % compared to the situation in 2016. the theoretical and practical implementation of the presented alternative objectives is the subject of further research [29]. in these more intensive considerations, embodied emissions are explicitly taken into account. with overall declining emissions, this part is becoming increasingly important. references [1] bayerisches staatsministerium für wirtschaft, energie und technologie. fortschrittsbericht 2016/2017 zum umbau der energieversorgung bayerns, germany, 2018. 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[25] l. schwan, j. hahn, m. barton, et al. analysis and development of reference buildings for the energy transition in the thermal sector. in proceedings of the 32nd international conference on efficiency, cost, optimization, simulation and environmental impact of energy systems, p. 1741. 2019. isbn 9788361506515. [26] m. barton, j. hahn, l. schwan, c. schweigler. bayerischer wohngebäudebestand und energiebedarf für die wärmeversorgung: aktueller stand und prognose für das jahr 2050. in book of abstracts inuas konferenz 2021, pp. 32–34. 2021. [27] m. barton, l. schwan, j. hahn, et al. development of reference building models based on statistical data for the analysis of the building heat demand. iop conference series: earth and environmental science 588(2):022051, 2020. https://doi.org/10.1088/1755-1315/588/2/022051 [28] m. barton, j. hahn, l. schwan, c. schweigler. comparison and evaluation of statistical databased reference building models for singleand multi-family houses. in proceedings of the 33rd international conference on efficiency, cost, optimization, simulation and environmental impact of energy systems (ecos 2020), p. 1947. 2020. isbn 9781713814061. [29] cenergie – center for energy-efficient buildings and districts, munich university of applied sciences, research group building construction and building climatology. [2022-01-20]. https: //www.ar.hm.edu/organisationen/ansprechpartner_ 3/professoren_3/essig/essig_2.de.html 435 https://doi.org/10.1088/1755-1315/588/2/022051 https://www.ar.hm.edu/organisationen/ansprechpartner_3/professoren_3/essig/essig_2.de.html https://www.ar.hm.edu/organisationen/ansprechpartner_3/professoren_3/essig/essig_2.de.html https://www.ar.hm.edu/organisationen/ansprechpartner_3/professoren_3/essig/essig_2.de.html acta polytechnica ctu proceedings 38:429–435, 2022 1 introduction 2 methods 3 results and discussion 4 conclusions references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.37.0043 acta polytechnica ctu proceedings 37:43–47, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague determination of radionuclide impurities in pure beta radiation sources by semiconductor spectrometry pavel novotný czech technical university in prague, faculty of nuclear sciences and physical engineering, department of dosimetry and application of ionizing radiation, břehová 7, 115 19 prague, czech republic correspondence: pavel.novotny@fjfi.cvut.cz abstract. in many applications, the radionuclidic purity of a radioactive substance is a crucial property. it involves, for example, the production of radiopharmaceuticals or reference standards used in metrology. however, the identification of admixtures is not straightforward for certain nuclides. the paper presents a technique for the determination of radionuclide impurities in the mixtures of pure beta-emitting nuclides. the proposed method is based on the comparison of measured beta spectra with the detector responses to individual radionuclides. these are obtained using a monte carlo calculation performed in the geometry corresponding to the accurate description of a measuring assembly. keywords: beta spectrometry, monte carlo simulation, si(li) detector. 1. introduction beta-decaying radionuclides are used as sources of ionizing radiation in various industrial, medical, and scientific applications. in many of them, the key parameter of a radioactive substance is its radionuclidic purity, i.e. the absence of other radionuclides in the source. usually, most of the potential impurities can be determined by gamma spectrometry. however, this analytical tool is not applicable if both the nuclide of interest and an admixture are pure beta radiation sources, i.e. no gamma ray is emitted. in addition, the impurity may be an isotope of the same element as the main nuclide, which precludes the use of any chemical analytical method. the proposal of a method for the identification and quantification of radionuclide impurities in the mixtures of pure beta-emitting nuclides was one of the objectives of the recently completed metrobeta project, which was supported by the european metrology programme for innovation and research (empir) [1]. the research focused on the measurement of beta radionuclides was performed by several european national metrology institutes. one procedure for identifying radionuclides from a mixed beta spectrum has been proposed by [2]. the authors demonstrated the application of the spectrum stripping method for the deconvolution of spectra measured using a magnetic spectrometer. the spectra belonging to individual nuclides were described by a function defined by the formula as follows 8∑ i=1 ai sin(bie + ci), (1) where e stands for the energy of radiation. the equation (1) is empirical in nature, which means that it describes the spectra shape well from a mathematical point of view, but it has no deeper physical meaning. the parameters ai, bi, and ci were obtained by fitting the experimental data. the method was tested for the analysis of mixed spectra of three radionuclides 32p, 90y, and 106rh with different mixing ratios. the accuracy achieved in determining the radionuclide activity ratios was better than 2 %. a different approach to solving the discussed problem is offered by the following procedure [3]. the mixed beta spectrum measured is divided into several regions. considering a mixture of three radionuclides a, b, and c, the count rate ni in each spectrum region can be expressed as ni = ηai aa + η b i ab + η c i ac, (2) where the symbol η denotes a detection efficiency. the desired activity values a are calculated by minimizing the expression representing the variance using the least square method. the procedures employed for the analysis of mixed beta spectra in this study partially utilize the principles mentioned above. however, values obtained from a precise monte carlo simulation were used to express the response to each single radionuclide instead of an empirical function. more details are given in the following section of the paper. 2. instrumentation and methods 2.1. detector and monte carlo model beta spectra were measured at the czech metrology institute spectrometry laboratory using a spectrometric system equipped with a lithium-drifted silicon detector. the detector was thoroughly characterized; therefore, its monte carlo model prepared for the mcnpx code [4] was available. a detailed description of the detector, its characterization, and the 43 https://doi.org/10.14311/app.2022.37.0043 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en pavel novotný acta polytechnica ctu proceedings model validation have been published in [5]. the source term for beta spectra calculations was defined as a histogram of electron energies derived from the fermi theory of beta decay. the simulated detector response to a given radionuclide was obtained using a pulse-height tally (type f8). the computation time of the simulations was chosen to achieve the relative type a uncertainty of the histogram bin values at the level of tenths of a per cent. the efficiency of calculations was improved by the variance reduction techniques (i.e. cell-by-cell energy cutoff and source biasing). all studied samples were prepared by drop deposition [6]. the droplet of a radioactive solution1 was deposited on a 0.1 mm thick polyethylene terephthalate foil. the solvent was subsequently evaporated, resulting in the formation of a thin radioactive layer. 2.2. two-window method suppose that the measured beta spectrum is a combination of contributions from two radionuclides a and b which differ in their beta decay endpoint energy. as the first step of an analysis using this method, two windows (regions of interest) are defined in the spectrum as indicated in figure 1. events originating from the radioactive decay of both components of a mixture are registered in region no. 1, whereas the detector response in region no. 2 is caused only by the radionuclide with a higher endpoint energy. figure 1. illustration of the two-window method idea. the following relations hold for the number of pulses n1 in region no. 1 and n2 in region no. 2 n1 = aaηa1 t + abη b 1 t, n2 = abηb2 t, (3) where ηi stands for the detection efficiency in the i-th region of the spectrum for the radionuclide indicated by the superscript value and t is the live time of the measurement. by dividing the first expression in equation (3) by the second one, a relation giving the ratio of radionuclide a and b activities is obtained aa ab = ηb2 ηa1 · ( n1 n2 − ηb1 ηb2 ) . (4) 1the aqueous solutions of srcl2 (20 mg of srcl2 per litre + 3 g of hcl per litre), ycl3 (50 mg of ycl3 per litre + 3 g of hcl per litre), and h3po4 (50 mg of h3po4 per litre) were used to prepare the samples analysed in this study. the spectrum in figure 1 corresponds to the situation where the endpoint energy value of an impurity is higher than the main nuclide endpoint. with a logarithmic scale on the vertical axis, its presence in a sample can be easily detected. on the other hand, if the impurity has an endpoint lower than the main nuclide, its identification may appear less clear, as the effect on the spectrum shape may not be visually apparent. nevertheless, the procedure for the quantitative evaluation is identical for both variants. 2.3. spectrum stripping method an alternative approach to the analysis of pure beta radionuclide mixtures is the spectrum stripping method. generally, this technique is based on the successive subtraction of the library spectra corresponding to individual nuclides present in a sample from the measured spectral response. the normalized library spectrum is subtracted channel by channel from the sample spectrum. this procedure proceeds from higher energies (i.e. for beta spectra from nuclides with a higher endpoint value) to lower ones. when analysing a mixed beta spectrum using this method, one region of interest is defined at first. its boundaries are chosen as follows. considering the difficulties in modelling the detector response for lowenergy electrons, it is advisable to choose the lower limit from approximately 500 kev upwards. the upper limit is restricted from above by the lowest value of an endpoint energy in the radionuclide mixture. in the selected region of interest, the integral of pulses is determined for both the measured and reference (i.e. admixture-free) spectra. for binary mixtures, the difference between these two values represents the number of pulses in the measured spectrum caused by the presence of an impurity in the sample ni. the number of pulses corresponding to the main nuclide nm is equal to the integral obtained from the reference spectrum. the activity ratio is then given by the following relation ai am = niη m nmηi (5) with ηi and ηm being the detection efficiencies for radionuclides i and m respectively. 3. results both methods of spectra deconvolution described above in sections 2.2 and 2.3 were tested on data from the measurement of samples with varying amounts of 89sr and 90sr. this mixture is commonly formed during the production of 89sr radionuclide due to parasitic reactions [7]. the samples also contained 90y, which is a radioactive decay product of 90sr. as the parent nuclide has a much longer half-life than the daughter nuclide2, it can be assumed that secular (long-term) radioactive equilibrium between these 2the half-lives of radionuclides 90sr and 90y are 28.80(7) a and 2.6684(13) d respectively [8]. 44 vol. 37/2022 radionuclide impurities in pure beta radiation sources figure 2. beta spectra of 89sr samples containg different amount of 90sr/90y impurity (top 0.3 %, bottom 3.0 %). nuclides was established in the sample. therefore, the activity of 90sr can be determined by the measurement of 90y. compared to 89sr, the endpoint energy of 90y is higher, which is favourable for the spectrum shape analysis because the high-energy part (> 500 kev) of beta spectra is not significantly distorted by self-absorption [9]. 3.1. two-window method the beta spectra of six samples were recorded using the si(li) detector. an example of two of these spectra is shown in figure 2. the ratio of the 90sr/90y impurity activity to the 89sr main nuclide activity was calculated using formula (4). the detection efficiencies were obtained from the monte carlo model of the detector. the regions of interest were defined in the energy range3 of (700–1400) kev and (1500–1700) kev. the results obtained are summarized in table 1. the standard uncertainty of the measured values was calculated as the uncertainty of an indirectly measured quantity (see [10]). 3the endpoint energy of 90sr equals 549.5(14) kev [8]. it implies that the radionuclide does not contribute to the detector response in the selected regions of interest. declared [%] measured [%] 0.100(1) 0.15(1) 0.300(3) 0.35(2) 0.500(5) 0.52(2) 1.00(1) 1.04(3) 2.00(2) 2.07(4) 3.00(3) 3.08(5) table 1. ratio of the 90sr/90y impurity activity to the main nuclide 89sr activity (two-window method). a second set of samples was analysed in the same way. these samples were also composed of strontium isotopes; however, 89sr was the impurity in 90sr/90y in this case. the calculated activity ratios are summarized in table 2. declared [%] measured [%] 1.00(1) 1.1(5) 2.00(2) 2.1(5) 5.00(5) 5.4(5) 8.00(8) 8.7(5) table 2. ratio of the 89sr impurity activity to the main nuclide 90sr/90y activity (two-window method). 3.2. spectrum stripping method the spectrum stripping method was tested on four spectra corresponding to different concentrations of the 89sr impurity in 90sr/90y. an example of one measured spectrum compared with the response to a 90sr/90y source of an older production date (i.e. virtually free of impurities4) is shown in figure 3. the ratio of 89sr to 90sr activity was calculated according to formula (5), with 90sr determined via 90y (assuming radioactive equilibrium between these two nuclides). the evaluation was performed in the energy interval of 1000 to 1490 kev. a comparison of the measured values with the declared ones is shown in table 3. the standard uncertainty of the results was calculated in accordance with the law of uncertainty propagation for an indirectly measured quantity [10]. declared [%] measured [%] 1.00(1) 1.0(4) 2.00(2) 2.1(4) 5.00(5) 5.2(4) 8.00(8) 8.3(4) table 3. ratio of the 89sr impurity activity to the main nuclide 90sr/90y activity (spectrum stripping method). 4the half-life of 89sr is 50.57(3) d [11], whereas it is equal to 28.80(7) a for 90sr [8]. 45 pavel novotný acta polytechnica ctu proceedings figure 3. beta spectra of two 90sr/90y samples (with and without 89sr impurity). 4. discussion specifically, the spectra of mixtures consisting of 89sr and 90sr/90y nuclides in various ratios were studied. it is easier to analyse a mixed beta spectrum if the admixture has a higher endpoint energy value than the main nuclide. the presence of the impurity is well evident in the spectrum, even for low concentrations of the order of tenths of a per cent. the measured and the declared activity ratios of the impurity to the main nuclide agree within the uncertainty for values greater than 0.5 %. the results are slightly overestimated for the lower concentrations. on the other hand, if the admixture endpoint energy is lower compared to the main nuclide, the analysis is more complicated. if only because the effect of the impurity on the spectrum shape is not obvious at first sight. moreover, the typical achievable uncertainties are larger due to the subtraction of close values in equation (4), which is used for the evaluation via the two-window method. the alternative evaluation of the same data set using the spectrum stripping technique provided slightly more accurate results (in terms of both measurement trueness and measurement precision). either way, both mentioned methods allow the activity ratio to be determined only at a level of a few per cent under these conditions. therefore, the sensitivity is about ten times lower in comparison with the situation mentioned in the previous paragraph. both proposed methods of beta spectra analysis are comparable in terms of instrumentation requirements or the complexity of calculations. however, the twowindow method seems to be slightly more universal. the spectrum stripping method is not suitable when the impurity has a higher endpoint energy than the main nuclide because the response from the admixture in the high-energy part of the spectrum is low and its statistics is poor. this fact complicates the normalization of reference spectra, which is crucial for the correct calculation of the impurity concentration. the advantage of the proposed method for the analysis of mixed beta spectra lies in its simplicity in terms of instrumentation. it requires only the equipment that is commonly available in radiometric laboratories. however, the results and the measurement repeatability are strongly dependent on the quality of measured samples. it is a limiting factor for the analysis of samples containing radionuclides with an endpoint energy value lower than circa 500 kev prepared by the universal drop deposition technique. figure 4. beta spectra of two 32p samples of the same design. the above-mentioned issue is demonstrated by the spectra in figure 4. both studied samples contained the same radionuclide 32p free of impurities and were of the same design. nevertheless, the measured spectra differ significantly from each other in shape in the low-energy part. as a result of the crystallization, the radioactive layer formed is not homogeneous and its thickness varies locally [9]. therefore, the level of selfabsorption in the samples also varies, which influences the low-energy part of the beta spectra. generally, there are two approaches how to overcome this limitation. the first relies on a change in the technology of samples production to improve their properties in terms of measurement repeatability. the second is based on the detailed characterization of each sample prepared by drop deposition, which is used to improve the accuracy of the monte carlo calculation of reference spectra. as part of the research on beta spectra metrology, the suitability of high-resolution x-ray fluorescence 2d scanning for this purpose was studied in detail and the results obtained have been presented in [12]. 5. conclusion the experiments performed showed that the analysis of beta spectra shape is a technique suitable for the quantification of radionuclide impurities. the method is promising, especially for the assay of mixtures consisting of pure beta minus emitters that are isotopes of one chemical element, i.e. other analytical tools, 46 vol. 37/2022 radionuclide impurities in pure beta radiation sources such as gamma spectrometry or chemical analysis, are not applicable. the findings are of direct practical relevance, and the proposed method can be easily implemented in radiometric laboratories where the radionuclide purity of substances is determined. references [1] metrobeta – radionuclide beta spectra metrology. [2022-04-28], http://metrobeta-empir.eu/. [2] u. lee, j. w. bae, h. r. kim. multiple beta spectrum analysis based on spectrum fitting. journal of radioanalytical and nuclear chemistry 314(2):617–622, 2017. https://doi.org/10.1007/s10967-017-5409-5. [3] y. kameo, a. katayama, a. fujiwara, et al. rapid determination of 89sr and 90sr in radioactive waste using sr extraction disk and beta-ray spectrometer. journal of radioanalytical and nuclear chemistry 274(1):71–78, 2007. https://doi.org/10.1007/s10967-006-6905-1. [4] d. pelowitz, j. durkee, j. elson, et al. mcnpx user’s manual version 2.7.0, 2011. [5] p. novotný, p. dryák, j. šolc, et al. characterization of the si(li) detector for monte carlo calculations of beta spectra. journal of instrumentation 13(01):p01021–p01021, 2018. https://doi.org/10.1088/1748-0221/13/01/p01021. [6] g. sibbens, t. altzitzoglou. preparation of radioactive sources for radionuclide metrology. metrologia 44(4):s71–s78, 2007. https://doi.org/10.1088/0026-1394/44/4/s09. [7] y. toporov, v. filimonov, r. kuznetsov, et al. production of 89sr using (n,γ)-and (n,p) reactions. in isotope production and applications in the 21st century, pp. 371–374. world scientific, 2000. https://doi.org/10.1142/9789812793867_0092. [8] m.-m. bé, v. chisté, c. dulieu, et al. table of radionuclides, vol. 3 of monographie bipm-5. bureau international des poids et mesures, sèvres, 2006. [9] k. siegbahn. alpha-, betaand gamma-ray spectroscopy. north-holland publishing co., amsterdam, 1965. p. 1367. [10] joint committee for guides in metrology. evaluation of measurement data – guide to the expression of uncertainty in measurement, 2008. [2022-04-28], https://www.bipm.org/utils/common/ documents/jcgm/jcgm_100_2008_e.pdf. [11] m.-m. bé, v. chisté, c. dulieu, et al. table of radionuclides, vol. 1 of monographie bipm-5. bureau international des poids et mesures, sèvres, 2004. [12] p. novotný. x-ray fluorescence analysis of samples for beta spectrometry and its application to precise monte carlo simulations. radiation physics and chemistry 192:109930, 2022. https: //doi.org/10.1016/j.radphyschem.2021.109930. 47 http://metrobeta-empir.eu/ https://doi.org/10.1007/s10967-017-5409-5 https://doi.org/10.1007/s10967-006-6905-1 https://doi.org/10.1088/1748-0221/13/01/p01021 https://doi.org/10.1088/0026-1394/44/4/s09 https://doi.org/10.1142/9789812793867_0092 https://www.bipm.org/utils/common/documents/jcgm/jcgm_100_2008_e.pdf https://www.bipm.org/utils/common/documents/jcgm/jcgm_100_2008_e.pdf https://doi.org/10.1016/j.radphyschem.2021.109930 https://doi.org/10.1016/j.radphyschem.2021.109930 acta polytechnica ctu proceedings 37:43–47, 2022 1 introduction 2 instrumentation and methods 2.1 detector and monte carlo model 2.2 two-window method 2.3 spectrum stripping method 3 results 3.1 two-window method 3.2 spectrum stripping method 4 discussion 5 conclusion references my title https://doi.org/10.14311/app.2022.36.0084 acta polytechnica ctu proceedings 36:84–89, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague maximizing information obtained from standardized fire tests using a bayesian framework balša jovanovića, b, ∗, ruben van coilea, edwin reyndersb, geert lombaertb, robby caspeelea a ghent university, department of structural engineering and building materials, technologiepark-zwijnaarde 60, 9052, ghent, belgium b katholieke universiteit leuven, department of civil engineering, kasteelpark arenberg 40, 3001 leuven, belgium ∗ corresponding author: balsa.jovanovic@ugent.be abstract. current practice is mostly focused on prescriptive design approaches where the performance of the structure in case of fire is assessed based on its performance in standardized fire tests. those tests indicate whether the structural member can withstand standardized iso 834 fire exposure for a certain code specified time. this method however does not provide an explicit safety level. this issue is enlarged even more by the fact that the iso 834 fire exposure does not represent a natural fire exposure, but a pseudo-worst-case exposure, making the correlation between the standardized fire test results and real-life behaviour of structural members exposed to fire questionable. however, there has been a century-old tradition of standardized fire tests with a lot of experience and infrastructure based on it. for that reason, here, a methodology is presented to obtain more information on the behaviour of structural members exposed to a natural fire from the standardized fire test results by using a bayesian framework. as an example structure, a simply supported concrete slab is considered. its failure during the standardized fire test is modelled and the parameters affecting the time when it fails (i.e., parameters affecting the nominal fire resistance time) are determined. the model is then used in a markov chain monte carlo procedure to update parameter distributions based on the measured fire resistance time. using these updated distributions, a full probabilistic calculation of the performance of the slab considering a natural fire exposure is then conducted to assess the failure probability. keywords: bayesian analysis, concrete, fire, safety level, uncertainty. 1. introduction structural stability of a building in case of fire is of utmost importance both for the safety of its occupants as well as for other fire safety objectives such as property protection. therefore, it is essential to provide a building design with an adequate safety level. in the structural sense, it is important to mitigate the risk of a collapse due to the fire effects. a common way of representing a structure’s performance in the case of fire is through its fire-resistance rating. the fire-resistance rating represents how long a structural member can withstand a furnace test before collapsing. during the furnace test, the structural member is exposed to a standardized fire exposure, a universal temperature-time curve that in its essence should represent a conservative worst-case fire scenario. this approach has a lot of benefits: it is standardized, it provides the same temperature exposure for a range of different materials and structural members, and therefore can be used to compare different solutions. however, it also has a few shortcomings. first of all, the standardized temperature-time curve does not represent any real fire scenario. this temperaturetime curve is monotonically increasing and therefore does not include the cooling phase of the fire. secondly, even though it should represent an envelope case for multiple possible fire scenarios, there are cases where this approach is not conservative [1]. despite these shortcomings, the standardized approach is highly popular. there is an abundance of data and research focused on the material and structural behaviour of members exposed to the standard temperature-time curve. unfortunately, the test does not provide any direct information about how the structure will behave in a real fire situation. all it provides is a time of failure for a structural member exposed to a fire scenario that, as mentioned earlier, might be conservative, but does not represent a real fire. this problem is even enhanced by the fact that a lot of current building codes allow the use of this value as proof of safety by prescribing the minimal fire resistance rating a member must achieve in the standardized test. in order to overcome this issue, in this paper, a bayesian updating procedure is applied. the bayesian framework in fire safety engineering has been successfully used in the past in a variety of fire safety problems. for example, wang and zabaras [2] used them in order to determine the magnitude of a heat source from the temperature data, guo et al. [3] used a zone 84 https://doi.org/10.14311/app.2022.36.0084 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 36/2022 standardized fire tests using a bayesian framework figure 1. temperature field of a 20 cm thick concrete slab with the concrete cover of 35 mm after 120 min of standardized fire exposure. model as a forward model for a bayesian inversion to obtain the fire size and its origin within a multicompartment structure based on the measured gas temperature. overholt and ezekoye [4] managed to invert the fire’s heat release rate based on the measured temperature data from room-scale experiments. the bayesian updating is used to predict the behaviour in a more realistic fire for the case of a simply supported concrete slab, considering the fire resistance as measured in the standardized test. using numerical heat transfer analysis, a bayesian framework and full probabilistic calculation, the failure time determined during the furnace test is used to compute the failure probability in case of a natural fire. 2. methodology 2.1. overview the procedure presented in this paper consists of two main parts. first, the prior distributions of the material properties of the slab are updated considering the measured fire resistance time. second, these posterior distributions are used to determine the failure probability of the slab in case of the natural fire. for the updating procedure, a model that calculates the failure time during the furnace test must be created. that is done by calculating the time instant when the capacity of the slab is lower than the loads applied. afterwards, this model is used in the bayesian framework to update the material properties based on the test failure time. using the updated material properties based on the test failure time the failure probability is calculated for the slab exposed to a natural fire. a monte carlo calculation is conducted taking into account uncertainties in the fire exposure, loads, geometry and updated material properties. finally, the failure probability is calculated for different values of test failure times and compared. 2.2. thermal calculation in order to determine the capacity of the concrete slab in the case of fire, the first step is to determine the temperate distribution inside of it. this requires the solution of a heat transfer problem which is commonly solved using the finite difference method. in this study, a 20 cm thick reinforced concrete slab exposed to the heating from the bottom side is considered. a fine mesh with the size of 0.1 mm and a time step of 0.1 s is used. the boundary condition on the bottom (heated) side is defined using a gas temperature-time curve and heat transfer through convection (convective heat coefficient hconv = 25 w/(m2.k) and radiation (emissivity ε = 0.7). on the top (cold) side, an ambient temperature of tamb = 20◦c is imposed. the density of concrete in the function of temperature is defined by en 1992-1-2 [5] with a moisture value of 3 %. a common approach of calculating for a concrete only section and assuming that the reinforcement temperature is equal to the concrete at the same position is employed. figure 1 shows the temperature distribution of the slab after 120 min of iso 834 fire exposure. the uncertainty associated with the parameters in the heat transfer problem is the focus of this paper. thus, probabilistic models for the thermal properties, thermal conductivity k and specific heat cp of concrete, are used. the models’ origin and details are presented in [6] and the main characteristics are summarized here. the main assumptions are that each of these properties is a continuous function of temperature, and at each temperature the values are distributed according to the gamma distribution. in order to obtain a random curve, the same quantile is used for all temperatures. for example, if the specific heat has a 5 % quantile value at 20 ◦c, it will also have a 5 % quantile value at 100 ◦c, 200 ◦c and so on. this way it is possible to define one continuous function for a thermal property at any temperature based only on the quantile value considered. this enables a quite easy implementation in numerical models and its use is presented in the following paragraph. first, a quantile value for each of the thermal properties is chosen independently, next mean and standard deviation at each needed temperature are calculated using the following formula (temperature t in ◦c), 85 b. jovanović, r. v. coile, e. reynders et al. acta polytechnica ctu proceedings parameter symbol unit distribution mean standard deviation ref. thermal conductivity k [w/(m.k)] gamma temperature dependent temperature dependent [6] specific heat cp [kj/(kg.k)] gamma temperature dependent temperature dependent [6] reinforcement yield strength fy [mpa] lognormal 560 30 [7] table 1. prior distribution for the parameters used in the bayesian framework. valid in the range 20 − 1200 ◦c [6]: kmean [w/(m.k)] = 6.627 · 10−7 · t 2 − 1.458 · 10−3 · t + 1.772 cp,mean [kj/(kg · k] = −2.953 · 10−7 · t 2 + 6.498 · 10−4 · t + 0.872 kstd [w/(m · k)] = 3.19 · 10−7 · t 2 − 0.691 · 10−3 · t + 0.434 cp, std [kj/(kg · k)] = −3.500 · 10−7 · t 2 + 7.700 · 10−4 · t + 0.042 (1) afterwards, using these values and given quantile, well-known equations defining the gamma distribution are used to calculate both thermal properties for each needed temperature. finally, the same moisture peak should be added to the specific heat curve as defined in en 1992-1-2 [5]. 2.3. bending capacity and failure time once the temperature inside the slab is calculated, its bending capacity can be calculated using the following formula [11]: mr = as · kf y (t ) · fy · ( h − c − ϕ 2 ) − (as · kf y(t ) · fy ) 2 2 · b · kf c(t ) · fc (2) where as is the reinforcement area, kf y (t ) is the temperature-dependent strength retention coefficient for the yield strength of steel, fy is the yield strength of steel in normal design conditions, h is the slab height, c is the concrete cover, ϕ is the reinforcement diameter, b is the width of the slab, kf c(t ) is temperature-dependent strength retention coefficient for the compressive strength of concrete and fc is the compressive strength of concrete in normal design conditions. in order to calculate when the slab would fail during the furnace fire resistance test, first the temperature distribution needs to be calculated for each point in time of the test. that is done using the described numerical heat transfer model and the iso 834 fire curve as boundary temperature-time curve. then, using equation 2 the slab’s capacity can be calculated for each point in time and from it consecutively the failure time can be calculated by determining the point in time when the capacity is lower than the applied loads. the loads during the test are assumed to be equal to 50 % of the slab’s capacity in normal conditions, (i.e., 50 % utilization in normal design conditions). this process is computationally expensive for later probabilistic calculations. hence, for 10 000 different combinations of the stochastic variables (table 1), a failure time is calculated and a polynomial regression model is fitted on the results (with the coefficient of determination being r2 = 0.99). all the other parameters in equation 2 have been regarded as constants instead of variables. 2.4. updating of uncertain parameters based on standardized fire test the updating procedure is a standard bayesian procedure using markov chain monte carlo simulations. the parameters that are updated are the quantiles of the thermal properties’ models and the yield strength of the reinforcement. all the priors are presented in table 1. the likelihood is defined under the assumption that the failure time error is normally distributed with a mean of 0 and standard deviation σerr and is presented in equation 3: l (y|x ) = 1 σ exp ( − 1 2 (y − m (x ))2 σ2err ) (3) where y is the measured failure time, m (x ) a model output for the input parameters x = (k, cp, fy ). the measurement uncertainty σerr is modelled by a normal distribution with a mean of 2 min and a standard deviation of 1 min, these values are chosen based on the model error and measurement precision during the furnace test. 2.5. natural fire exposure the end goal is to obtain the probability of failure in the case of a real fire using the measured test failure time. for that, the thermal gradient for a natural fire exposure also has to be calculated. a parametric 86 vol. 36/2022 standardized fire tests using a bayesian framework parameter symbol unit distribution mean standard deviation ref fuel load gf [mj/m2] gumbel 780 234 [8] concrete cover c [mm] beta [µ ± 3σ] 40 5 [9] dead load g [knm] normal 27 2.7 [10] live load (qk = 16) q [knm] gamma 3.2 3.04 [10] load model uncertainty ke [−] lognormal 1 0.1 [10] resistance model uncertainty kr [−] lognormal 1 0.15 [10] table 2. distributions of the parameters used to calculate the failure probability. figure 2. posterior distributions for thermal conductivity (left) and specific heat of concrete (right) at different temperatures for different values of failure time with the shaded 90 % highest density intervals (hdi). fire curve from en 1991-1-2 [8] is used to represent a natural fire exposure. in order to use it, some parameters defining the compartment where the slab is located must be defined. in this study, a compartment with dimensions of 10 m × 10 m × 3 m, opening factor o = 0.05m1/2 and wall linings with thermal inertia of 1450 j/(m2ks1/2) is considered. these parameters can be considered as easily determined and constant. however, to define the parametric fire curve, the fuel load in the compartment also needs to be defined. there is a lot of uncertainty on this parameter and therefore it will be considered stochastically (see table 2). similarly, as with the analysis of the furnace test, a heat transfer calculation is needed to determine the temperature gradient and the slab’s bending capacity. again, polynomial regression is used as there is a need for a large number of calculations. for 1000 different combinations of the thermal conductivity quantile, the specific heat quantile and the compartment fuel load, the temperature gradient is calculated and that data is used for the regression. the final step is the calculation of the conditional failure probability given a structurally significant fire. this is done using the updated material properties’ distributions and additional stochastic parameters (presented in table 2). the failure probability is calculated using equation 4 which describes the limit state z. z = kr · mr − ke · (g + q) (4) 3. results and discussion the bayesian updating was performed multiple times for different assumed values of the failure time during the test. updated posterior probability density functions of both thermal conductivity and specific heat at different temperatures for a few failure times are presented in figure 2. from these results, it is observed that the failure time is not particularly sensitive to the value of the thermal conductivity and the updating procedure results in only small changes to the prior distribution. for example, at ambient temperature, the standard deviation reduces from 0.42 to approximately 0.29 w/(m.k). however, the failure time is highly dependent on the value of the specific heat. the standard deviation at ambient temperature reduces from 0.57 for the prior to roughly 0.13 kj/(kg.k) for the posterior. furthermore, an almost linear relation between the posterior mean value of the specific heat and the failure time can be seen, as shown in figure 3 (left). using the first bayesian method (mcmc) the yield strength of reinforcement steel was also updated, how87 b. jovanović, r. v. coile, e. reynders et al. acta polytechnica ctu proceedings figure 3. posterior mean value of the specific heat of concrete at ambient temperature in function of the failure time (left); posterior distributions of the reinforcement steel yield strength at ambient temperature for different values of the failure time (right). figure 4. failure probability (left) and reliability index (right) in the function of the failure time. ever, according to the results in figure 3 (right), it is evident that the failure time sensitivity to this parameter is relatively low. it must be mentioned that the mcmc sampling method might produce unrealistically low posterior uncertainty as in theory the update would only be specifically valid for the tested slab and not account for the possible variations between similar slabs that are going to be used for the same structure (which are not tested). in the case where more than one test is used, a procedure similar to the one presented in [12] should be used where the mechanical properties of reinforcement are not updated in order to focus on the thermal properties. to determine how much the failure probability changes with the updated posteriors, the failure probability is first determined based on the prior distributions. this is calculated using 106 monte carlo realizations and the obtained value is equal to 4.4 · 10−3. . the probability of failure is then also calculated using the posterior distributions using the posteriors for thermal conductivity and specific heat instead of their respective priors. the results are presented in figure 4 in the function of the failure time. in order to demonstrate how different test failure times affect the probability of failure in the case of the real fire, the failure probability has been calculated for the range of test failure times from 110 to 150 min and it ranges from 8.35 − 2.03 · 10−3. as expected, a higher value of the failure time leads to a lower value of the failure probability. however, what is interesting is that for failure times lower than 88 vol. 36/2022 standardized fire tests using a bayesian framework 120 min, the failure probability is higher than the one calculated using the prior distributions. when analysing the distributions of the moment capacity of the slab, it is observed that the reduction of the variability compared to using prior distributions is almost negligible. this can be attributed to the fact that the uncertainty on the resistance of the slab is mainly governed by the uncertainty on the fuel load. however, the posteriors of the thermal properties still have an effect, as it is observed that the mean value of the capacity changes for different measured failure times. 4. conclusions a procedure to obtain a more precise failure probability of a structural member in case of fire using bayesian updating based on the standard furnace test result was presented. this procedure focuses on maximizing information from the highly used furnace test in relation to the performance-based assessment of the structural capacity under a real fire condition. as the furnace test provides the failure time in case of standardized fire exposure, a clearer connection from its results to the structure’s behaviour in real fire is explored. the procedure is demonstrated on the case of a reinforced concrete slab. a standard markov chain monte carlo simulation approach was used and showed that the failure time during the test is most sensitive to the variation of the specific heat. for example, the updating procedure resulted in a reduction of the standard deviation of specific heat at ambient temperature from the prior value of 0.57 to a posterior value of approximately 0.13 kj/(kg.k). this also has an effect on the calculated failure probability under a real fire condition. using the updating procedure and monte carlo simulations, the failure probability under a real fire condition was calculated for the range of test failure times from 110 to 150 min and it ranges from 8.35 − 2.03 · 10−3. it was also observed that the variability of the concrete slab capacity in the case of a natural fire is mainly determined by the uncertainty of the fire exposure and that a reduction of the uncertainty of material properties obtained with this method has a small but distinctive effect. it must be noted that this study’s results are limited to the behaviour of the reinforced concrete slab, where the failure is mostly a consequence of strength reduction of the reinforcement due to the elevated temperature. in the cases of other materials and structural members, the effectiveness of this approach might differ. nevertheless, this presents an improvement in the use of the furnace test results as this approach tackles one of its biggest flaws, correlation to the real fire behaviour, and shows that they can be used to quantitatively assess the safety level of a structure in case of a real fire. acknowledgements the authors wish to thank the research foundation of flanders (fwo) for the financial support on the research project (grant number 3g010220) "vibration-based postfire assessment of concrete structures using bayesian updating techniques". references [1] a. law, l. bisby. the rise and rise of fire resistance. fire safety journal 116, 2020. https://doi.org/10.1016/j.firesaf.2020.103188. [2] j. wang, n. zabaras. a bayesian inference approach to the inverse heat conduction problem. international journal of heat and mass transfer 47(17-18):3927-41, 2004. https://doi.org/10.1016/j. ijheatmasstransfer.2004.02.028. [3] s. guo, r. yang, h. zhang, et al. new inverse model for detecting fire-source location and intensity. journal of thermophysics and heat transfer 24(4):74555, 2010. https://doi.org/10.2514/1.46513. [4] k. j. overholt, o. a. ezekoye. characterizing heat release rates using an inverse fire modeling technique. fire technology 48(4):893-909, 2012. https://doi.org/10.1007/s10694-011-0250-9. [5] cen, en 1992-1-2:2004: eurocode 2: design of concrete structures part 1-2: general rules. structural fire design. european standard, 2004. [6] b. jovanović, n. e. khorasani, t. thienpont, et al. probabilistic models for thermal properties of concrete. proceedings of the 11th international conference on structures in fire (sif2020), p. 342-352, 2020. https://doi.org/10.14264/363ff91. [7] jcss, jcss probabilistic model code. part 3.2 static properties of reinforcing steel. joint committee on structural safety, 2001. [8] cen, en 1991-1-2:2002 actions on structures part 1-2: general actions actions on structures exposed to fire. european standard, 2002. [9] jcss, probabilistic model code. part 3.10 dimensions. joint committee on structural safety, 2001. [10] b. jovanović, r. van coile, d. hopkin, et al. review of current practice in probabilistic structural fire engineering: permanent and live load modelling. fire technology 57(1):1-30, 2020. https://doi.org/10.1007/s10694-020-01005-w. [11] t. thienpont, r. van coile, r. caspeele, et al. burnout resistance of concrete slabs: probabilistic assessment and global resistance factor calibration. fire safety journal 119, 2021. https://doi.org/10.1016/j.firesaf.2020.103242. [12] w. botte, e. vereecken, l. taerwe, et al. assessment of posttensioned concrete beams from the 1940s: large-scale load testing, numerical analysis and bayesian assessment of prestressing losses. structural concrete 22(3):1500-22, 2021. https://doi.org/10.1002/suco.202000774. 89 https://doi.org/10.1016/j.firesaf.2020.103188 https://doi.org/10.1016/j.ijheatmasstransfer.2004.02.028 https://doi.org/10.1016/j.ijheatmasstransfer.2004.02.028 https://doi.org/10.2514/1.46513 https://doi.org/10.1007/s10694-011-0250-9 https://doi.org/10.14264/363ff91 https://doi.org/10.1007/s10694-020-01005-w https://doi.org/10.1016/j.firesaf.2020.103242 https://doi.org/10.1002/suco.202000774 acta polytechnica ctu proceedings 36:84–89, 2022 1 introduction 2 methodology 2.1 overview 2.2 thermal calculation 2.3 bending capacity and failure time 2.4 updating of uncertain parameters based on standardized fire test 2.5 natural fire exposure 3 results and discussion 4 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0354 acta polytechnica ctu proceedings 38:354–360, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague twin rooms – new experimental test cells for testing advanced facade elements boris bielek∗, daniel szabó, josip klem, kristína kaniková, alžbeta danková slovak university of technology in bratislava, faculty of civil engineering, radlinského 11, 810 05 bratislava, slovak republic ∗ corresponding author: boris.bielek@stuba.sk abstract. nowadays the global trend is the integration of new materials, constructions and technological principles, which are simultaneously implemented in individual scientific and engineering disciplines. reducing the energy intensity of buildings will increasingly resonate in individual political and professional circles. as a result, new fragments of building envelopes in the field of facade engineering are being developed and tested. testing of building envelope is carried out either in static (laboratory) boundary conditions or in dynamic (climatic, real) conditions. the determination of the test method is conditioned by the specific intention or the investigated phenomenon within the construction of the building envelope. currently, we finished the development and realization of the new experimental facility twin rooms for testing advanced elements of building envelopes in dynamic boundary conditions (in the real climate of central europe – bratislava) in terms of building thermal engineering and energy efficiency of buildings. it is based on the concept of pavilion measurement. the essence of the research is that the outdoor climate is modelled by the conditions of the real outdoor climate. test cells consists of a solar laboratory – two-room for a comparative study of the effect of solar radiation and heat transfer on energy consumption and indoor climate. the space of two identical laboratory rooms is situated inside a container – a pavilion, whose climate is a compensating space. only the tested facade element walls are exposed to the outdoor climate. the exchange of energy with the environment is possible only through this measured facade wall. the article brings a detailed description of this experimental equipment, basic technical parameters of its technological circuits and methodology of experimental measurements. keywords: experimental pavilion test cells, advanced facade elements, real outdoor climate. 1. introduction in the current period, humanity is solving one of the most difficult technical and economic problems of ensuring sufficient energy for continual development of society without negative impact on ecology, production and environmental protection. approximately in the 1980s, the economic system expressing human activities exceeded the reproductive value of the earth’s biocapacity (3 to 3.5 billion tons of emissions of greenhouse gases) and the man began to produce the ecological debt on the planet. the world, at the present time, produces environmental burden that exceeds its biocapacity by about 30 %. humanity must find a solution for the repayment of the ecological debt. for a successful solution to this problem, mankind has to reconsider its priorities in investments, which represent the transformation to a sustainable society by transforming the energy sector towards environmentally clean renewable energy sources and their conversion, transformation of material sector towards ecologically clean materials and their production and transformation of the whole economy towards lowenergy and low-emission technologies and environmentally friendly products. sustainability is synonymous with the image of a future world whose aim is to place man, nature and technology in permanently stable equilibrium. this means such a development which allows both current and future generations to meet their basic living needs, preserves the diversity of nature and functions of the ecosystems. from the internal structure of the overall energy balance of non-productive sphere, the human settlements are the second largest consumer of energy. the majority of this energy consumption falls on the operation of buildings (heating, cooling, ventilation, hot water and lighting), coupled with hygienic and ecological comfort in them. there are two major approaches to sustainable development of society in the interaction energy – human settlements: energy saving by its rational use and focus on environmentally clean renewable energy sources. renewable energy sources, as the dominant production technology value of nature, with its simultaneous renewal of ecosystems are becoming determining factors for changes in fundamental concept of energy quantification of buildings. building is becoming a place for the collection and on-site conversion of renewable energy as part of the transformation of organization of the energy market. it is becoming part 354 https://doi.org/10.14311/app.2022.38.0354 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 twin rooms – new experimental test cells . . . of the energy distribution networks and it is just in the interaction to these distribution networks that the new quantification of physical-energy demand of the building is being conceived, expressed by the term net zero energy building or nearly net zero energy buildings or net plus energy buildings. facade structures are indispensable and, by their functions, extremely important parts of the building envelopes. besides their physical (thermal, sound and water insulation and air impermeability) and mechanical features (static and dynamic wind resistance) they also have additional specific indispensable functions (natural daylight, natural ventilation, visual contact with the surroundings, fire escape route, architectural expression). with the development of technology in the architecture towards a sustainable society, the function of transparent structures is highlighted as an element of passive solar systems of buildings. given the need to regulate energy flows through transparent constructions within dynamically changing conditions of the external climate, new elements are being integrated to modern transparent building constructions (adjustable ventilation louvers, flaps or units, mobile screening devices...), connected to electronic sensors of physical or chemical parameters of internal or external climate that dynamically respond to climate changes and optimize the energy flows through transparent constructions in order to ensure a comfortable indoor climate while minimizing demands on building environmental technology (i.e. energy). with development of active solar systems based on photo thermal or photoelectric conversion, these elements are then integrated directly into transparent facade structures. these transparent structures then partially take over the function of building environmental technology. with the development of modern climate-adaptive transparent facades capable of dynamically responding to the climatic parameters of the external climate and thus optimizing the energy flows through their construction in order to significantly reduce the heat load in the summer period, the heat losses in the winter period and to maximize the available renewable energy sources, the need to verify their parameters and energy efficiency depending on the conditions of the external climate comes to the fore. 2. overview of experimental equipment for research of advanced facade elements research in the test cell began in 1985 as an effort to increase faith in the application of energy-conscious and passive solar building products and evaluation techniques used to provide practical thermal properties of the products. the passys (passive solar components and systems testing) project focused on test cell equipment as a means of determining the performance of passive solar building components and providing additional information on building design and simulation tools. the advantage of test cells is that they provide a well-controlled, realistic room (cell) environment without occupancy [1]. the passys test cell consists of a well-insulated construction 8 × 2.7 × 2.7 m with two rooms. one room measuring 5 × 2.7 × 2.7 m is a south-facing test cell and an adjacent north-facing area is a service room that contains an air conditioning unit. it is possible to test various building elements on the south facade of the test room. for this purpose, manuals for instruments, operations, calibrations and test procedures have been developed. the cells are able to test vertical and horizontal building elements. the passys test cells are located in european countries: belgium, denmark, france, germany, italy, the netherlands, scotland, greece, portugal and spain [2]. similar test cells in the rest of europe are vliet test building at ku leuven university in belgium [3, 4], test box at technische hochschule rosenheim in germany [5], minibat test cell in lyon, france [6], fact in cea-ines, le bourget du lac in france [7], zeb test cell laboratory in trondheim, norway [8, 9] and the cube in aalborg, denmark [10–12]. the vliet test building research facility is one of the full-scale test cells. the research is solved at the component level and focuses on the combined transfer of heat, air and moisture through walls, roofs and floor systems and their impact on sustainability. the methodology includes modelling, accelerated aging tests, hot-box measurements and “in situ” studies [3, 4]. the test box was provided to the technical university of rosenheim, germany, from the facade construction company josef gartner, for which it acquired all the rights to research, development and training. in addition to evaluating comfort and energy efficiency, they also evaluate the effect of shading, angle of blinds to obtain the most daylight with the least possible glare. therefore, the test cell is rotatable about a vertical axis. thus, it can be adapted to the requirements of the respective test program. of particular interest are comparative studies on energy efficiency, comfort and visual comfort in real conditions [5]. the minibat test cell in lyon, france, consists of two adjacent cells connected by a door located in a compensating container. one side of the first cell opens into a climatic chamber that models the outdoor climate. the walls and interior rooms are equipped with many sensors and both cells can be ventilated, heated, cooled, etc. the aim of this test device is to gain information about heat and air transfers within one room of a building, between two rooms and between a cell and the outdoor climate simulated by a weather generator. the device allows to collect detailed data for validation of numerical models. typical studies include air distribution (indoor air quality, analysis of airflow from inlet openings and their mixing with indoor air), heating, cooling (energy efficiency, humidity and heat comfort for different ventilation systems and heating / cooling elements) and facade systems (facade elements, 355 b. bielek, d. szabó, j. klem et al. acta polytechnica ctu proceedings double skin facades, phase change materials, etc.) [6]. fact (facade tool) is a test facility at cea-ines, le bourget du lac, in south-eastern france. it is an experimental facility consisting of a two-storey building with 10 rooms for testing the perimeter cladding of buildings with different widths and heights of facades, with different geometries of the internal environment. the size of the test cells can be modified so that measurements can be performed in one room or in open space [7]. the zeb test cell laboratory in trondheim, norway, is used to test low-energy integrated building systems with real operating conditions. the test cell can be divided into two smaller chambers, in which different technologies can be used to compare results. in the laboratory, various elements of building materials, building envelopes, energy installations and control systems are jointly developed and optimized in real climate conditions [8, 9]. the university of aalborg uses a test cell, namely the cube, which is one of the full-scale test facilities, for research double skin facades. it is a test cell situated in a compensation space with one side facing south, into which the test sample is inserted. the cell is designed to be able to adapt to operating modes, natural or mechanical flow conditions, various shading techniques and similar. the thermal conditions in the cell can be perfectly controlled, as well as in the room next to it. the test facility is equipped to allow the measurement of any power supplied to the experimental zone in order to maintain the required temperature conditions [10–12]. there are currently several testing facilities at slovak and czech universities in košice [13], žilina [14], brno [15–17] and prague [18]. the laboratories in košice include 3 outdoor experimental cells, which are designed for monitoring, research of physical properties and characteristics of facade structures – thermaltechnical and humidity problems in real conditions of use of buildings and real outdoor climatic conditions [13]. pavilion research has been built in the uniza research centre in žilina, the main part of which consists of two differently oriented (east and south) experimental walls. stable indoor environment parameters are ensured in the test cells. temperature and relative humidity sensors are located in the experimental walls. in the middle of the test cells there are sensors for measuring temperature and humidity. outdoor climate conditions are obtained from a meteorological station located on the roof of the building [14]. in the research centre centrum admas (advanced materials, structures and techniques) in brno, four small test cells were built in 2014 to research the thermal stability of buildings. the exterior walls of the cells have different material bases (brick, aerated concrete and wooden construction). sensors are placed in the cells to measure surface temperature, air temperature, to quantify the amount of solar radiation in the planes on the surface of the glazing. the outdoor climate is recorded by their meteorological station’s prototype. the research centre also includes a mobile experimental full-scale test set of solar-activated facades and their concepts [15–17]. the university centre for energy efficient buildings uceeb at the czech technical university in prague has a climate facade with a controlled indoor environment and 6 monitored fields with samples exposed to the weather [18]. however, it must be stated that none of the above-described test facilities in the territory of slovakia and the czech republic makes it possible to accurately quantify the energy flows of the examined sample. 3. “twin rooms” construction the “twin rooms” test cells are used to determine the thermal properties (u-value, thermal resistance, surface temperatures, temperature fields) and energy flows through advanced facade elements in the range of outdoor climatic conditions. the technological equipment for the “twin rooms” must ensure controllable, stable, homogeneous and reproducible indoor temperature conditions in the two measuring cells and in the compensation room. this will make it possible to perform measurements of heat balances, total and local heat flows as well as surface temperatures simultaneously on two tested structures installed in measuring openings under the influence of non-stationary outdoor climatic conditions. the experimental test cells consist of an officetype mobile container with external dimensions of 6058 × 2990 × 4200 mm, which is placed on pre-built concrete foundations. the supporting structure of the container itself consists of a space frame welded from steel hollow rolled profiles. the envelope consists of lightweight sandwich panels, with an external galvanized shaped sheet, a filling high-efficiency thermal insulation based on mineral wool and an internal lining made of chipboard. the entrance to the container is situated on the rear longer side of the container (6058 × 4200 mm) and is made by a door opening with dimensions of 1000 × 2000 mm with galvanized doors with infill thermal insulation – figure 1. the south-facing front side of the 6058 × 4200 mm container has two holes cut out with 2400 × 3400 mm, within which the test element of the building envelope will be fitted. directly behind the openings, a dividing structure of a light wooden partition creates two rooms (twin rooms) with dimensions of 2400×2000×3400 mm thermally insulated from the interior space of the container – figure 1. 4. “twin rooms” technology the technology of the test cells “twin rooms” consists of several parts, which form one functional unit. using environmental technology, the same constant indoor climate θai [°c] will be maintained in both cells and the compensating space of the container, which will ensure zero heat flow through the dividing structures separating the two cells from the inner space 356 vol. 38/2022 twin rooms – new experimental test cells . . . figure 1. construction schemes of an experimental test cells twin rooms for testing advanced facade elements in real conditions of outdoor climate. of the container. by ensuring this, all energy flows will be only through the tested facade elements of the building envelopes fitted in the measuring openings of the container. in order to maintain a constant temperature θai [°c] in the two measuring cells, we must supply energy (heat or cold) to those cells. the amount of this supplied energy is identical to the energy that was released by the measured facade element. we can measure the energy delivered to individual cells (rooms) and thus we can also measure the energy that was released by through the facade element. in this way we can measure all energy flows through facade elements in conditions of dynamically changing outdoor climate over time. the technological equipment for air conditioning of the compensation room (container) consists of an outdoor air conditioning unit exterior air – room air for the room “container” with a volume of about 25 m3 with an output of about 3 kw and a heating element with an output of 2 kw – figure 2. this ensures a stable temperature in the compensation room to prevent the exchange of thermal energy between the measuring cells and the compensation room. at the same time, it is necessary to ensure even air circulation through the circulating fans in the compensation room, in order to avoid local temperature inhomogeneities of the air surrounding the measuring cells. technological equipment for air conditioning of testing cells (rooms) 1 and 2 consists of air conditioning unit (identical as for container) air exterior – water exchanger for space of cells 1 and 2 with a volume of about 15 m3 with an output of 6 kw and a heater with an output of 2 kw – figure 2. this ensures a stable and homogeneous temperature in the cells. in this case, it is necessary to continuously measure the power supplied to the air-conditioned space of cell, so on the air-conditioned side there is a water exchanger measuring water flow and its inlet and outlet temperature as well as an electricity meter to measure the electrical power of the heater and circulating fans. to evaluate the total heat fluxes by the measured structures, it is necessary to measure parallel in “cell 1” and in “cell 2”: • the amount of energy supplied or taken by the heat exchanger (inlet and outlet water temperature – media flowing through the exchanger, the volume of water flowed through the exchanger), • the amount of electrical energy supplied by the radiator and circulating fans. guaranteed parameters for container space: • cooling circuit power 3 kw, • heat circuit power 2 kw, • adjustable temperature range: ▷ for summer period + 19 °c to + 40 °c, homogeneity in space and time ± 1.5 °c, ▷ for winter period + 10 °c to + 25 °c, homogeneity in space and time ± 1.5 °c, • air velocity range: 0.10 m/s to 1.50 m/s. 357 b. bielek, d. szabó, j. klem et al. acta polytechnica ctu proceedings figure 2. construction scheme of technological equipment of an experimental test cells twin rooms. 1 – cooling source, qc = 16.4 kw, mex ea 117 z c, 2 – plate heat exchanger – part of the supply of the cooling source, 3 – water tank – storage tank galmet 140 l, 4 – pipe fancoil nxkt3-v1000, 5 – sill fancoil nxkh3-v600, 6 – circulating fans – homogeneous thermal field, 7 – electrical switchgear, 8 – opening for mounting a new facade sample. guaranteed parameters for cell space 1 and cell space 2: • cooling circuit power 6 kw, • heat circuit power 2 kw, • adjustable temperature range: ▷ for summer period + 19 °c to + 40 °c, homogeneity in the space in front of the measured sample ± 1.5 °c (at the output of the cold and heat circuit up to 2.0 kw), ▷ for winter period + 10 °c to + 25 °c, homogeneity in the space in front of the measured sample ± 1.5 °c (at the output of the cold and heat circuit up to 2.0 kw), • air velocity range: 0.10 m/s to 1.00 m/s. all sensed physical quantities will be continuously recorded and evaluated by the local control unit. the system enables remote access in real time to all sensed and evaluated quantities as well as remote control of the entire control and regulation system. 5. “twin rooms” measuring equipment in order to be able to evaluate the total heat balance of the measured structures under the influence of nonstationary outdoor climatic conditions, it is necessary to measure these conditions with the help of a meteorological station. the weather station will be located on the roof of the container with measuring cells. the temperature and humidity of the outside air, the speed and direction of the wind, the intensity of sunlight on the vertical and horizontal planes and the amount of precipitation will be measured continuously. to evaluate the local thermal parameters of the measured structures, it is necessary to measure: • local heat fluxes using heat plates, • surface temperatures on the outside and inside of the structure by pt100 sensors, • flow velocities along the measured structures, • air pressure differences on the outside and inside of 358 vol. 38/2022 twin rooms – new experimental test cells . . . the structure, • intensity of solar radiation on the vertical surface of the perimeter wall. all the above-mentioned physical quantities are continuously measured and stored by means of a measuring control panel. at the same time, they are archived by means of a pc and are remotely accessible at any place via the internet. 6. initial measurements on the new experimental test cells twin rooms after the successful delivery of technology to the new twin rooms test cells for research of advanced facade elements in real climate conditions (october 2021) and after its successful test operation, we plan to start experimental verification of a prototype of a new modular double skin transparent facade with a narrow slot cavity with a usage of multi-stage renewable energy of solar radiation, which was developed within the scientific projects apvv-16-0126 and vega 1/0113/19 in cooperation with ingsteel s.r.o. bratislava. the resulting design of the double facade is characterized by: • progressive construction of the completed spatial part – an element mounted from the exterior to the elements of the supporting system of the building, characterized by fast assembly technology and the removal of seasonality, • high thermal-technical quantification without thermal bridges of frame profiles based on aluminium and anchoring systems and high acoustic quantification of the facade, • application of the design solution of a double skin transparent facade with a narrow slot cavity and the height of the section identical to the height of one floor, which significantly eliminates the load from solar radiation in summer and at the same time reduces heat losses during the heating period – winter, • application of the ventilation heat recovery unit in the parapet part of the element, which in winter uses preheated air from the physical cavity of the double skin facade with a positive impact on a significant reduction of heat losses of the building from ventilation, • application of photovoltaics for direct conversion of solar radiation into electrical energy in the parapet part of the outer transparent wall of the double skin facade for direct use in the building, • application of an automated control system of individual structural elements of the element (closing of inlet and outlet openings, movement of shading devices, control of heat recovery unit, etc.) based on measuring outdoor climate parameters and indoor climate requirements. the resulting energy flows of the new double skin transparent facade will be compared with the classic simple transparent facade of similar physical quantification as the inner wall of the double facade, also with the application of the same ventilation heat recovery unit and photovoltaics in its parapet part. we will test both samples of facades by long-term research in the experimental twin rooms test cells during the period of 1 year. 7. conclusion after the construction of the experimental test cells twin rooms for pavilion research of advanced facade elements in conditions of real outdoor climate, the facility will enable research in the field of modern facade technology of buildings using available renewable energy sources (especially solar radiation) with integrated passive and active systems and also implementation of development and optimization of structural creation of modern transparent facade structures for realization sphere. acknowledgements this research was supported by scientific grant agency mšvvš sr and sav under vega 1/0113/19 and by the 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[2021-09-29]. https://www.web-old.uceeb.cz/system/files/ souboryredakce/laboratore/experimentalnifasada/rp1_experimentalni-fasada.pdf 360 https://www.zeb.no/index.php/en/test-cell-laboratory https://www.zeb.no/index.php/en/test-cell-laboratory https://doi.org/10.1016/j.egypro.2017.09.718 https://doi.org/10.1016/j.buildenv.2014.09.009 https://doi.org/10.4028/www.scientific.net/amm.861.213 https://doi.org/10.4028/www.scientific.net/amm.861.213 https://doi.org/10.4028/www.scientific.net/amm.824.519 https://doi.org/10.4028/www.scientific.net/amm.824.519 https://www.web-old.uceeb.cz/system/files/souboryredakce/laboratore/experimentalni-fasada/rp1_experimentalni-fasada.pdf https://www.web-old.uceeb.cz/system/files/souboryredakce/laboratore/experimentalni-fasada/rp1_experimentalni-fasada.pdf https://www.web-old.uceeb.cz/system/files/souboryredakce/laboratore/experimentalni-fasada/rp1_experimentalni-fasada.pdf acta polytechnica ctu proceedings 38:354–360, 2022 1 introduction 2 overview of experimental equipment for research of advanced facade elements 3 ``twin rooms'' construction 4 ``twin rooms'' technology 5 ``twin rooms'' measuring equipment 6 initial measurements on the new experimental test cells twin rooms 7 conclusion acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.36.0216 acta polytechnica ctu proceedings 36:216–223, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague a study on the effects of hidden safety when assessing existing structures max teichgräber∗, daniel straub technical university münich, school of engineering and design, engineering risk analysis grou, theresienstraße 90, 80333 münchen, germany ∗ corresponding author: max.teichgraeber@tum.de abstract. in many instances, the safety of existing structures can no longer be demonstrated by standard code-based assessments. reasons for this include changes in the code, changes in the demands on the structures and deterioration. to address this problem, it is common practice to perform a more detailed assessment utilizing advanced models. in this way, many structures can be shown to comply with safety requirements, even if they cannot be verified by standard assessments. the standard code models are often conservatively biased. this leads to designs which include hidden safety. if the reassessment is performed with more advanced models in lieu of standard models, the hidden safety can vanish. concurrently, the reduced uncertainty of advanced models may compensate this safety reduction. in this paper we investigate this issue on a hypothetical population of existing bridge structures under traffic. we consider that the standard code model is exchanged by an advanced traffic load simulation. keywords: code based design, existing bridge structures, hidden safety, traffic load. 1. introduction structural design codes aim to provide easy-to-use rules that lead to economical and safe structures [1]. today’s structural design codes are based mainly on the semi-probabilistic partial safety factor (psf) concept [2–5]. it contains two explicit safety components: the psfs and the characteristic values. both components are determined via code calibration [6–8]. the calibration is typically based on standard models and the determined safety components are intended to be used with these standard models. in many cases the standard models are biased in a conservative sense (e.g. [9]). this leads to another implicit safety component: the hidden safety. the presence of hidden safety is unproblematic, as long as standard models are applied; however, problems may arise if standard models are exchanged by more advanced models. thereby, two effects are present: first, advanced models are typically not biased; hence, the hidden safety vanishes. second, advanced models typically include less model uncertainty. when it comes to the design of new structures, these two effects influence the structural reliability in opposite directions. if one effect dominates, a recalibration of the explicit safety components of the psf concept is needed. the recalibration ensures that, on average, the standard and the advanced models lead to the same level of safety. we investigate this in detail in [9, 10]. in the assessment of existing structures both effects are present as well. the difference to new designs is that for existing structures the design choices are already made; hence, the advanced models are only used to reassess if a given design is in compliance with the requirements. whether or not the accepted design is safe enough is the question addressed in this paper. we perform a case study on a hypothetical population of existing bridge structures under traffic load. we consider an increase of the traffic load since the construction date of the bridges such that they are no longer in compliance with standard code-based assessments. we investigate how the application of advanced traffic load simulation affects the reliability. 2. design following the partial safety factor concept we follow the approach and the nomenclature of the eurocode [11]. the investigations can be transferred to other semi-probabilistic structural design codes. in a eurocode design, four different models can be identified (figure 1): the load model ml,ec , the structural model ms,ec , the material model mm,ec , and the resistance model mr,ec . we use the subscript ec to stress that these are the standard models provided by eurocode. the load model and the material model are typically probabilistic models. in order to make the design process deterministic, a characteristic load lk,ec and characteristic material properties mk,ec are determined. to consider the uncertainty of the load model and the material model and to ensure a sufficiently safe structure, the characteristic values are modified by the psfs γf and γm. the resulting values are the input of functions ts,ec and tr,ec provided by the structural model and the resistance model. to consider the uncertainty of these models, the outcome is 216 https://doi.org/10.14311/app.2022.36.0216 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 36/2022 a study on the effects of hidden safety load model ml,ec characteristic load lk,ec partial safety factor γf structural model ms,ec partial safety factor γsd design load effect ed,ec characteristic material mk,ec partial safety factor γm resistance model mr,ec partial safety factor γrd design resistance rd,ec material model mm,ec figure 1. overview of the eurocode design appraoch. again modified by psfs γsd and γrd. eventually, the design values ed,ec and rd,ec are obtained. a design is sufficient, if the design values fulfill the following inequality: ed,ec ≤ rd,ec (1) ⇔ γf · ts,ec (lk,ec · γsd) ≤ tr,ec ( mk,ec γm ) γrd for the sake of simplicity, the eurocode merges the partial safety factors of the action and the resistance side: γf = γf × γsd (2) γm = γm × γrd (3) in the subsequent case study we assume ts,ec and tr,ec to be linear functions through the origin. then equation 1 simplifies to: γf · ts,ec (lk,ec ) ≤ tr,ec (mk,ec ) γm (4) if the design is chosen optimally in regard to resource consumption, equality holds in equation 4. from this, pec is derived: pec = γf · γm · ts,ec (lk,ec ) tr,ec (mk,ec ) (5) pec can be interpreted as the design choice following the psf concept (e.g. the cross section area of a truss). the corresponding probability of failure can be determined via the following limit state function (lsf): g = pec · tr(m ) − ts (l) (6) note that in equation 6 tr, ts , m and l are not subscripted with ec, hence, do not originate from the eurocode. m and l are the "true" – or as we call them: purely aleatoric [9] – distributions of the load and the material property. tr and ts represent the true relationship between the load and the material property to the load effect and the resistance. generally, these are unknown and have to be estimated carefully. the probability of failure pr(f ) results from pr(f ) = ∫ ωf fm,l(m, l) dm dl (7) where fm,l(m, l) is the joint probability density function (pdf) of m and l and ωf = {m, l | g < 0} is the failure domain. 3. case study: existing bridge structures under traffic load in the following we perform a case study on a hypothetical population of existing bridge structures under traffic load. 3.1. setup and assumptions we choose the following setup of the case study: • n = 10 000: number of samples of the population of bridge structures. this is needed, since some of the subsequent calculations are sample-based. apart from the statistical error originating from the sampling, the results are not effected by n. • we choose the material property m to follow a lognormal distribution: m ∼ ln e[m ] = 1 c. o. v.[m ] = 0.1 (8) the coefficient of variation of 0.1 is in the typical range of material properties (e.g. concrete compression strength) [12]. some material properties may have higher (e.g. timber bending strength) or lower coefficients of variation (e.g. structural steel yielding strength). therefore, we vary the coefficient of variation between 0.05 and 0.3 in a sensitivity analysis. • we choose the annual maximum of the traffic load l to follow a standardized gumbel distribution: l ∼ g e[l] = 1 c. o. v.[l] = 0.08 (9) the choice is based on simulated traffic following the approach of [13]: local traffic data from three locations of german highways with light, medium and heavy traffic is analyzed. on this bases, traffic streams (sequences of vehicles) are generated applying numerical traffic simulation. the simulated traffic load is applied on a portfolio of bridges. the portfolio includes bridges with 1-3 spans, each with 10, 15, or 20 m. on each bridge we apply either light, medium and heavy traffic. hence, the portfolio consists of 3 · 3 · 3 = 27 bridges. for each bridge a structural analysis is carried out and the resulting 100-year time histories of the internal forces are evaluated at all locations that are needed for the design. the annual block maxima of the time histories are taken to fit generalized extreme values distributions via the maximum-likelihood-method. the shape parameters of the fitted generalized extreme value distributions are close to 0, in which case the 217 max teichgräber, daniel straub acta polytechnica ctu proceedings generalized extreme value distribution converges to the gumbel distribution. the sample mean of the coefficient of variation of the field moments is 0.06, of the supporting moments 0.08 and of the shear forces at the supports 0.08. this justifies the choice of c. o. v[l] = 0.08. note that l is the purely aleatoric distribution; hence, c. o. v[l] = 0.08 only covers aleatoric uncertainties. epistemic uncertainties are added via the error in the estimation of the characteristic load (see θec and θadv below). we alter the coefficient of variation between 0.05 and 0.3 in a subsequent sensitivity analysis. a coefficient of variation of 0.3 is rather high in the context of traffic loads. by this, we ensure to also cover design situations with high uncertainty (e.g. due to rerouting of roads). • as already mentioned, we assume that the functions ts,ec (translates the loads to the load effects) and tr,ec (translates the load property to the resistance) are both linear and trough the origin. we further assume that both functions are equal to the true functions ts and tr. with this assumptions the functions ts,ec , ts , tr,ec and tr cancel each other out within the lsf (equation 6). • we assume a traffic load increase since the construction of the bridge by a factor of finc = 1.3. this value is not on the bases of any data but should be interpreted as a hypothetical increase. in an actual application case, this value must be derived respectively. we vary finc between 1 and 2 in a subsequent sensitivity analysis. • following the traffic load model of the eurocode [14], we define the characteristic traffic load as the 99.9% quantile of l (1000 year return period which corresponds to the probability of exceedance of 5% in 50 years). we model the error in the estimation of the characteristic value via a relative error: θadv = lk lk,adv (10) θec = lk lk,ec (11) where lk is the characteristic value of the purely aleatoric traffic load, lk,adv is the estimation of lk by means of advanced traffic load modeling and lk,ec is the estimation of lk following the eurocode. we choose θadv and θec to follow lognormal distributions: θadv ∼ ln e[θadv ] = 1 c. o. v[θadv ] = 0.1 (12) θec ∼ ln e[θec ] = 0.7 c. o. v[θec ] = 0.2 (13) the choices of the distribution-parameters are justified as follows: ▷ e[θadv ] = 1: since the simulation is based on real data, it is reasonable to assume that it estimates the characteristic value without a bias. ▷ c. o. v[θadv ] = 0.1: the coefficient of variation comprises three different uncertainties: the statistical uncertainty due to a limited simulation time, the uncertainty of the chosen distribution type and the uncertainty of incomplete modeling (e.g. construction sites on the highway may change the traffic load, but are not included in the simulation). the statistical uncertainty is estimated via the multivariate normal distribution of the parameter estimates of the fitted extreme value distributions of the inner forces [15]. the resulting coefficients of variation due to statistical uncertainty are in the range of 0.02-0.04. we – unfortunately – lack any data to estimate the uncertainty of the two other sources; hence, the overall uncertainty of c. o. v[θadv ] = 0.1 is based on our subjective assessment. we alter c. o. v[θadv ] between 0.05 and 0.2 in a subsequent sensitivity analysis. ▷ e[θec ] = 0.7: within the portfolio of bridge structures the sample mean of the relative error is 0.76 regarding field moments, 0.70 regarding supporting moments and 0.79 regarding the shear forces at the supports. we apply a value of 0.7 and alter e[θec ] = between 0.4 and 1.0 in a subsequent sensitivity analysis. ▷ c. o. v[θec ] = 0.2: θec can be rewritten as θec = lklk,adv · lk,adv lk,ec = θadv · lk,adv lk,ec . the distribution parameters of θadv is already known/estimated. the distribution of lk,adv lk,ec is also lognormal. the mean coefficient of variation of lk,adv lk,ec is 0.17 regarding field moments, 0.13 regarding supporting moments and 0.14 regarding the shear forces at the supports. we follow the biggest value of 0.17. this results in a a coefficient of variation of θec of approximately 0.2. we alter c. o. v[θec ] between 0.1 and 0.3 in a subsequent sensitivity analysis. • following eurocode [11], we define the characteristic material property as the 5% quantile of m . we assume that mk,ec exactly meets mk, hence, that no estimation error in the determination of the characteristic value is present. • following eurocode [11] we choose the psf γf = 1.35. we further set γm = 1.1. 3.2. numerical investigations our main objective is to investigate the probability of failure of the bridges within the hypothetical population designed by eurocode and the change of the probability of failure when more advanced traffic load models are used. we investigate how parameter changes in the setup influence the mean probability of failure and the acceptance rate of reassessed bridge structures. our secondary objective is to understand how the probability of failure would change if the advances models would have been used to design the bridges. 218 vol. 36/2022 a study on the effects of hidden safety we only cover this issue for the setup defined in section 3.1 and do not investigate parameter changes within the setup. by rearrange equation 10 and 11, we determine the distribution of the characteristic traffic load according to eurocode and according to advanced traffic load modeling by solving equation 11 and 10 for lk,ec and lk,adv . the resulting distributions are: lk,ec ∼ ln µln lk,ec = ln(lk) − µln θec σln lk,ec = σln θec (14) lk,adv ∼ ln µln lk,adv = ln(finc · lk) − µln θadv σln lk,adv = σln θec (15) where µln and σln are the location and the scale parameter of the respective lognormal distribution. we sample lk,ec,i and lk,adv,i from lk,ec and lk,adv simulating the estimation of the characteristic traffic load of i = 1, ..., 10 000 bridges. we sample from lk,ec and lk,adv independently. given one sample lk,i (either from the characteristic value of the eurocode or the advanced traffic load model) the probability of failure is calculated as: pr(f ) = ∫ ωl fm ( finc · l · mk γm · γf · lk,i ) · finc · fl(l) dl (16) first, we calculate the probability of failure for the case of finc = 1. from this our secondary objective (how the probability of failure changes if the advances models would have been used to design the bridge) can be studied. in a next step, we increase the traffic load finc > 1 to investigate how the increase in the load changes the probability of failure and how a reassessment using advanced traffic load modeling affects the probability of failure). the reassessment decision is based on the following train of thought: advanced traffic load modeling may estimate a characteristic traffic load that is lower than the characteristic traffic load according to eurocode (even when an increased traffic load is present). if this lower characteristic value would be used to design the bridge, it would result in smaller resistances; therefore, the reassessment according to advanced traffic load modeling would deem the existing eurocode design to be sufficient. from this, the following acceptance/rejection rule is derived: lk,adv,i ≤ lk,ec,i ⇒ accept (17) lk,adv,i > lk,ec,i ⇒ reject (18) 3.3. results and discussion figure 2 shows the probabilities of failure (annual reference period) of the bridges designed by eurocode loaded with the original traffic load (finc = 1). the mean annual probability of failure is e[pr(f | ec-design, finc = 1)] = 4.9 · 10−8. this probability can be interpreted as the target probability of failure, 10−30 10−25 10−20 10−15 10−10 10−5 100 0 200 400 600 800 e[pr(f | ec-design, finc = 1)] = 4.9 · 10−8 figure 2. histogram of the probability of failure pr(f | ec-design, finc = 1) of bridge structures designed following eurocode. since the original design criteria reaches this probability of failure on average and these designs were accepted by society. the probability is rather low compared to the common range of probabilities of failures of structures or compared to the target probability of failure defined in the eurocode of 8.5 · 10−6. the reasons for the deviation are as follows: we do not consider model uncertainties of the structural model, the material model and the resistance model. moreover, we applied only traffic load on the bridge structures. if also other loads would be considered (with a higher coefficient of variation compared to the coefficient of variation of the traffic load), the probability of failure might be higher. we do these simplifications to keep the issue simple. this way the effects hidden safety are more isolated and more straightforward to study. in case of an actual application these simplifications should not be conducted. due to these simplifications the absolute value of the probability of failure may not be very meaningful. however, a relative comparison to the probabilities of failure of the bridges designed or accepted/rejected by advanced traffic load models is valid, since the probabilities of failure are calculated on the same basis. figure 3 shows the probabilities of failure that would result if the bridges (loaded with the original traffic load finc = 1) would have been designed using advanced traffic load modeling. the resulting mean probability of failure e[pr(f | adv-design, finc = 1)] = 3.1 · 10−7 is below the target of 4.9 · 10−8. this shows, that the negative effect on the structural reliability of the lost hidden safety due to advanced traffic load modeling is stronger than the positive effect of the reduced model uncertainty. an additional safety factor would be needed if advanced traffic load models were to be used for bridge design. here, we do not conduct a calibration of such a factor. a detailed guide is 219 max teichgräber, daniel straub acta polytechnica ctu proceedings 10−30 10−25 10−20 10−15 10−10 10−5 100 0 200 400 600 800 e[pr(f | adv-design, finc = 1)] = 3.1 · 10−7 figure 3. histogram of the probability of failure pr(f | adv-design, finc = 1) of bridge structures designed following advanced traffic load modeling. 10−30 10−25 10−20 10−15 10−10 10−5 100 0 200 400 600 800 e[pr(f | ec-design, finc = 1.3)] = 1.1 · 10−5 figure 4. histogram of the probability of failure pr(f | ec-design, finc = 1.3) of bridge structures designed following eurocode loaded by a 30 % higher traffic load. given [9]. remark: the resulting probabilities of failure of the advanced designs with an increased load (finc > 1) would be exactly the same as in figure 3, as long as the load increase is also considered within the design. finc would appear twice and cancel each other out: once within the design (a factor applied on lk,adv,i) and once as a factor applied on the load l. figure 4 shows how the probabilities of failure of the bridges designed according to eurocode change if the traffic load is increased by 30 % (finc = 1.3). the increased traffic load raises the mean probability of failure from 4.9 · 10−8 to 1.1 · 10−5. figure 5 subdivides the histogram of figure 4 into accepted/rejected cases (according to advanced traffic load modeling). 68.8 % of the bridge structures are accepted and 31.1 % are rejected. the mean probability of failure of the accepted bridges 10−30 10−25 10−20 10−15 10−10 10−5 100 0 200 400 600 800 e[pr(f | ec-design, finc = 1.3), accepted] e[pr(f | ec-design, finc = 1.3), rejected] = 3.5 · 10−5 = 1.7 · 10−8 figure 5. histogram of the probability of failure pr(f | ec-design, finc = 1.3) of bridge structures designed following eurocode loaded by an 30 % increased traffic load divided into accepted (green) and rejected (red) bridges according to advanced traffic load modeling. e[pr(f | ec-design, finc = 1.3), accepted] = 1.7 · 10−8 is smaller than the target probability of failure of 4.87 · 10−8; hence, the reassessment leads to sufficiently safe structures. therefore, the use of advanced models for reassessment of existing structures appears justified without further adjustment of the safety factors. this is in contrast to the design of new structures. the deifference can be understood by comparing the histogram of the probability of failure of the accepted bridge structures (green histogram in figure 5) to the histogram of the probability of failure of bridge structures designed by advanced traffic load modeling (blue histogram in figure 3): following equation 17 a structure is accepted if lk,adv ≤ lk,ec . if equality would holds, the accepted eurocode design and the advanced design are equal, hence both histograms would coincide. the greater the difference lk,ec − lk,adv , the more hidden safety remains, which decreases the probability of failure in the reassessment case. we conclude that the application of advanced models to reassess a given structure with increased load • retains parts of the hidden safety of accepted structures; • decreases the probability of failure by reducing uncertainty. figure 6 shows the changes in the mean probability of failure and the acceptance rate changes, when altering selected parameters of the case study one at a time. the reassessment of bridge structures should be critically considered, if the mean probability of failure of the accepted structures is greater than the mean 220 vol. 36/2022 a study on the effects of hidden safety 1.0 1.1 1.2 1.3 1.4 1.5 1.6 1.7 1.8 1.9 2.0 10−4 100 10−2 10−10 10−12 10−8 10−6 finc 1.0 1.1 1.2 1.3 1.4 1.5 1.6 1.7 1.8 1.9 2.0 0.8 1.0 0.2 0 0.4 0.6 finc 0.4 0.5 0.6 0.7 0.8 0.9 1.0 10−4 100 10−2 10−10 10−12 10−8 10−6 e[θec ] 0.8 1.0 0.2 0 0.4 0.6 e[θec ] 0.10 0.15 0.20 0.25 0.30 10−4 100 10−2 10−10 10−12 10−8 10−6 c. o. v.[θec ] 0.8 1.0 0.2 0 0.4 0.6 c. o. v.[θec ] 10−4 100 10−2 10−10 10−12 10−8 10−6 c. o. v.[θadv ] 0.8 1.0 0.2 0 0.4 0.6 c. o. v.[θadv ] 0.05 0.10 0.15 0.20 0.25 0.30 10−4 100 10−2 10−10 10−12 10−8 10−6 c. o. v.[m ] 0.05 0.10 0.15 0.20 0.25 0.30 0.8 1.0 0.2 0 0.4 0.6 c. o. v.[m ] 0.05 0.10 0.15 0.20 0.25 0.30 10−4 100 10−2 10−10 10−12 10−8 10−6 c. o. v.[l] 0.05 0.10 0.15 0.20 0.25 0.30 0.8 1.0 0.2 0 0.4 0.6 c. o. v.[l] 0.4 0.5 0.6 0.7 0.8 0.9 1.0 0.10 0.15 0.20 0.25 0.30 0.05 0.10 0.15 0.20 0.10 0.15 0.20 0.25 0.30 figure 6. mean probability of failure (left) and acceptance rate (right) of bridge structures altering setup parameters one at a time. the blue lines represent the mean probability of failure when the original load is applied (target probability of failure), the teal dashed lines represent the mean probability of failure when the increased load is applied and the dash doted green and red lines represent the mean probability of failure of the accepted and the rejected bridge structures. 221 max teichgräber, daniel straub acta polytechnica ctu proceedings probability of failure of the bridge structures under the original load (target probability of failure). this is the case if • the load strongly increases (finc > 1.6). however, this case is not critical, since the target probability of failure is not exceeded by much and the acceptance rate simultaneously drops very low. • the coefficient of variation of the material property is rather large (c. o. v[m ] > 0.2). this is also not critical, since the target probability of failure is only exceeded marginally. • the coefficient of variation of the traffic load is rather large (c. o. v[m ] > 0.15). again, this is not critical, since the target probability of failure is only exceeded marginally. • the estimation of the characteristic traffic load according to eurocode is very biased (e[θec ] < 0.65). this is also not concerning, since in this case the eurocode design includes high amounts of hidden safety, which leads to very low probabilities of failure. this hidden safety can serve as a reserve against increased traffic loads. • the estimation of the characteristic traffic load according to eurocode is associated with low uncertainty (c. o. v[θec ] < 0.18). in this case the coefficient of variation in the estimation of the characteristic value of the eurocode model and the advanced traffic load model match closely. hence, the reassessment does not reduce the uncertainty significantly, but reduces the bias (hidden safety). this case can be critical and should be prevented; however, this case is rare, since the uncertainty of advanced models is typically much smaller than the uncertainty of standard models. • the estimation of the characteristic traffic load according to advanced models is relatively uncertain (c. o. v[θadv ] > 0.18). as in the previous bullet point the coefficient of variation in the estimation of the characteristic value of the eurocode model and the advanced traffic load model are close. this case can also be critical and should be prevented; however, again is rather rare. overall, the reassessment of bridge trough advanced modeling in most cases leads to sufficiently safe structures. it is only critical, if c. o. v[θec ] and c. o. v[θadv ] are very similar. however, the conduced sensitivity analysis only varies one parameter at a time. a case in which multiple parameters are varied simultaneously may lead to a more critical outcome. 4. conclusion we investigated the effects of hidden safety if advanced traffic load models replace the standard eurocode traffic load model. in the design case, the loss of hidden safety increases the probability of failure more than the reduced uncertainty of advanced models decreases it; therefore, advanced traffic load models should not be applied to design new structures without recalibrating the partial safety factors. the reassessment case showed less critical results. in the vast majority of considered cases the accepted bridge structures that are assessed as compliant are sufficiently safe and no recalibration is needed. in this case the hidden safety is only reduced proportional to the load increase and not fully erased (in contrast to the design case). references [1] o. ditlevsen, h. madsen. structural reliability methods. wiley new york, 1996. [2] b. ellingwood, t. galambos, j. macgregor. development of a probability based load criterion for american national standard a58: building code requirements for minimum design loads in buildings and other structures. department of commerce, national bureau of standards, 1980. [3] t. galambos, b. ellingwood, j. macgregor, a. cornell. probability based load criteria: assessment of current design practice. journal of the structural division 108(5):959–977, 1982. https://doi.org/10.1061/jsdeag.0005958. [4] b. ellingwood, j. macgregor, t. galambos, a. cornell. probability based load criteria: load factors and load combinations. journal of the structural division 108(5):978–997, 1982. https://doi.org/10.1061/jsdeag.0005959. [5] deutsches institut für normung. grusibau: grundlagen zur festlegung von sicherheitsanforderungen für bauliche anlagen [grusibau: basic principles for the definition of safety requirements for structures]. beuth verlag, 1981. [6] h. madsen, s. krenk, n. c. lind. methods of structural safety. courier corporation, 2006. [7] m. faber, j. sørensen. reliability-based code calibration: the jcss approach. proceedings of the 9th international conference on applications of statistics and probability in civil engineering 9:927–935, 2003. [8] m. baravalle. risk and reliability based calibration of structural design codes. ph.d. thesis, norwegian university of science and technology, 2017. [9] m. teichgräber, j. köhler, d. straub. hidden safety in structural design codes. engineering structures 257:114017, 2022. https://doi.org/10.1016/j.engstruct.2022.114017. [10] m. teichgräber, j. köhler, d. straub. über den umgang mit versteckten sicherheiten eine fallstudie am windlastmodell des eurocode [on dealing with hidden safeties a case study on the wind load model of the eurocode]. baustatik – baupraxis 14 2020. [11] cen. eurocode 0: basis of structural design, 2002. [12] joint committee on structural safety. the jcss probabilistic model code. 2001. [13] m. nowak, o. fischer. site-specific load models for road bridges – an important tool for advanced evaluation strategies in bridge reassessment. betonund stahlbetonbau 112(12):804–814. https://doi.org/10.1002/best.201700064. 222 https://doi.org/10.1061/jsdeag.0005958 https://doi.org/10.1061/jsdeag.0005959 https://doi.org/10.1016/j.engstruct.2022.114017 https://doi.org/10.1002/best.201700064 vol. 36/2022 a study on the effects of hidden safety [14] cen. eurocode 1: actions on structures, 2005. [15] s. coles, j. bawa, l. trenner, p. dorazio. an introduction to statistical modeling of extreme values. springer, 2001. 223 acta polytechnica ctu proceedings 36:216–223, 2022 1 introduction 2 design following the partial safety factor concept 3 case study: existing bridge structures under traffic load 3.1 setup and assumptions 3.2 numerical investigations 3.3 results and discussion 4 conclusion references my title https://doi.org/10.14311/app.2022.36.0015 acta polytechnica ctu proceedings 36:15–24, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague structural assessment and expected utility gain sebastian thönsa, b a lund university, faculty of engineering, box 117, 221 00 lund, sweden b bam federal institute for materials research and testing, unter den eichen 87, 12205 berlin, germany correspondence: sebastian.thons@kstr.lth.se abstract. this paper contains an introduction, probabilistic formulation, and exemplification (1) of system state and utility actions, (2) system state and utility action value analysis and (3) a threshold formulation for a predicted information and action decision analysis. the approaches build upon structural condition assessment and provide a basis for the condition management by maximising the expected utility for information and actions before information acquirement and action implementation. following the basic distinction of system state and utility actions, strengthening, replacement, repair, load reduction and consequence reduction actions are formulated. with an exemplary study encompassing an expected utility calculation and an action value analysis, it is demonstrated how expected utility optimal physical system changes can be identified before implementation. the threshold formulation for a predicted information and action decision analysis relies on the equality of the decision theoretical posterior action optimality condition. by extending the exemplary predicted action decision analyses with different structural health information (shi), the threshold formulation is exemplified and the optimal condition management strategies are identified. keywords: structural condition management, action uncertainty modelling, action and information value. 1. introduction condition assessment is usually understood as a structural reliability analysis with additional information. most commonly it is performed with obtained information, i.e., by posterior structural reliability updating. however, condition assessment constitutes a part of the structural condition management, i.e., ensuring the integrity and functionality of a system over time. the structural condition management may require measurements to learn about the structural condition and actions to adapt the structural condition. these two basic means of structural condition management should be performed efficiently, i.e., in a way that measurements and actions are planned to maximise their technological, economic and risk reduction effects. an efficient structural condition management can be performed and planned by the utilisation, adaptation and the application of the bayesian decision theory [1] to structures; or more widely to built environment systems. the decision theoretical context of structural condition assessment and management necessitates (1) a system state model, (2) a utility model, (3) that the information acquirement state is considered in the probabilistic modelling and (4) that the relevant actions can be modelled. whereas points (1) to (3) have been extensively developed over the last decade, see e.g., [2–17], the action modelling remains a challenge. relevant actions for a condition management would encompass e.g., structural modifications and utility changes. further, it would be required to analyse the expected action utility prior to the decision to implement an action. this paper focusses on the introduction and formulation of system state and utility actions as well as the action value analysis building upon [18] and [19] in sections 2 and 3. the action modelling in section 2 includes the introduction of an action uncertainty modelling and the formulation of strengthening, replacement, repair, load reduction and utility actions in conjunction with a limit state function representative for a highly correlated series system including extreme events and deterioration. based on the action uncertainty modelling, predicted and implemented action decision analyses are formulated (section 3). in section 4, the action modelling is exemplified and expected utility theorem consistent action values for various actions are quantified. section 5 contains a decision theoretical formulation of predicted information and action decision analyses (pipa da) and a derivation for continuous and direct structural health information (shi). it is shown how this shi formulation can be discretised and solved. the in section 4 introduced example is further developed in section 6 and the value of the condition management is quantified and separated into measurement information and system state and utility actions. section 7 includes and a summary and concluding remarks. 15 https://doi.org/10.14311/app.2022.36.0015 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en sebastian thöns acta polytechnica ctu proceedings 2. system performance and action modelling this section encompasses the formulation of the system state analysis, i.e., a structural reliability analysis in a decision theoretical context, and the combination of system state reliability and a utility model to quantify the system performance (section 2.1). in section 2.2, a probabilistic action modelling in introduced. 2.1. system state analysis and system performance analysis a generic component limit state function gx1 , describing the component state event x1, may be formulated with the resistance r, the damage d and the loading s, which are subjected to the resistance and loading model uncertainties mr, md and ms , respectively: x1 : gx1 = mr · r (md · d) − ms · s ≤ 0 (1) the resistance may be decomposed in the section property z and the material strength rmat. both may be subjected to the influence of deterioration, i.e., the design variable is subjected to section deterioration (md,z · dz ) and the material strength is influenced by degrading material properties (md,rmat · drmat ): x1 : gx1 = mr · z (md,z · dz ) · rmat (md,rmat · drmat ) − ms · s ≤ 0 (2) in assessing normatively an expected utility-based decision or a decision value, the benefits, costs and consequences of the system states are modelled. based on these models, the preference of decision makers can be described using the concept of a utility function implying the context of a descriptive decision analysis [20]. in engineering decision analysis, the utility function is rather linear and usually based on monetary values, where the utilities of the system states are expressed in terms of costs or, generally, negative consequences. in the following, the utility function is thus omitted and a normative decision analysis performed. the system states are described by xl (see above) resulting in a utility u (xl) . the system performance usp , i.e., the expected system utility, is then calculated as the expected value of a probability mass function (where the system states denote realisations of a discrete random variable), i.e.: usp = exl [u (xl)] = ∑ xl u(xl) · p (xl) (3) 2.2. action modelling the expected system performance (see above) can be influenced by physical changes affecting (1) the utility of the system states and (2) the resistance and/or the load of the system states. this basic distinction leads to the denotation of utility actions and system state actions, respectively. the latter type can be associated with actions performed by an engineer as it includes e.g., design, repair and strengthening actions. beyond the distinction of action types, an action implementation state and uncertainty modelling is required to analyse and to optimise the expected of the utility before action implementation (see section 1). it is thus in the further distinguished between a predicted or implemented action. the explicit modelling of the m implementation states yk,m and the modelling of implementation uncertainties are introduced. a utility action is described with the system state utilities u(xl, ak, yk,m), the implementation states yk,m and the implementation uncertainties p (yk,m) and the costs of the action c(ak). the index m may be used for distinguishing e.g., with m = 1 . . . 3 full, partial or failed action implementation. a system state action ak is described with the system state probabilities p (xl(ak, yk,m)) beside the implementation uncertainties and the costs. in the context of existing structures, examples for system state actions are e.g., repair, replacement, strengthening and loading actions. a repair action can be modelled by reversing deterioration effects due to mechanical reshaping of the component. the repair action including the action implementation uncertainties (ak, yk,m) is then allocated to the damage, i.e., d(ak, yk,m) affecting the material resistance or the section properties: x1 : gx1 = mr · z (md,z · dz ) · rmat (md,rmat · drmat (ak, yk,m)) − ms · s ≤ 0 (4) x1 : gx1 = mr · z (md,z · dz (ak, yk,m)) · rmat (md,rmat · drmat ) − ms · s ≤ 0 (5) a replacement action would result in a damage free component with same, similar or new material properties: x1 : gx1 = mr · z · rmat(ak, yk,m) − ms · s ≤ 0 (6) a strengthening action results in an enhanced resistance of a component by e.g., section enlargement, material addition or the replacement with a stronger material, i.e., z(as,k, yk,m, . . .) and rmat(as,k, yk,m, . . .): 16 vol. 36/2022 structural assessment and expected utility gain utilitiesactions system state action chance chancechoice system state and utility management implementationsystem states kalx utilities u (objective) function ( ) lsp x l u e u x= é ùë û ,k my utility action ( ) ( ) ( )s k ,m l k k ,mkpa y l kx a ,ya u maxe e u x c aé ù= -é ùë ûë û ( ) ( ) u k ,m l k pa y x l k,m k ka u maxe e u x ,y ,a c aé ùé ù= -ë ûë û system statesactions ka lx implementation ,k my utility u system performance (sp) analysis predicted system state action (psa) decision analysis predicted utility action (pua) decision analysis figure 1. system performance functions and objective functions for predicted action decision analyses. x1 : gx1 = mr · z(ak, yk,m, md,z · dz ) · rmat(md,rmat · drmat ) − ms · s ≤ 0 (7) x1 : gx1 = mr · z(md,z · dz ) · rmat(ak, yk,m, md,rmat · drmat ) − ms · s ≤ 0 (8) actions on the load side can be a load application for testing purposes or the change of use leading to a modified loading, s(as,k, yk,m): gx1(as,k ,yk,m) = mr · z(md,z · dz ) · rmat(md,rmat · drmat ) − ms · s(ak, yk,m) ≤ 0 (9) utility actions may include a benefit enhancement due to an enhanced functionality or a consequence reduction in case of failure. a utility action is subjected to action implementation uncertainties dependent on the system states, i.e., u(xl, yk,m, ak). for illustration purposes, the expected utility quantification of an implemented action analysis can be written as: uia = eyk,m [exl [u (xl, yk,m, ak) − c(ak)]] = ∑ xl ∑ yk,m p (xl) · p (yk,m) · u(xl, ak) (10) a utility action can be further detailed to consequence reduction ((c(xl, ak), equation (11)), benefit and functionality enhancement (b(xl, ak) , equation (12)) and (equation (13)) cost reduction actions in conjunction with the (normative) utility model characteristics: uia = ∑ xl ∑ yk,m p (xl) · (b(xl) − p (yk,m) · c(xl, ak) − c(ak)) (11) uia = ∑ xl ∑ yk,m p (xl) · (p (yk,m) · b(xl, ak) − c(xl) − c(ak)) (12) uia = ∑ xl ∑ yk,m p (xl) · (b(xl) − c(xl) − p (yk,m) · c(ak)) (13) 3. predicted action decision analysis and action value analysis the objective function for a predicted action decision analysis (pa da) can be written for system state and utility actions. for both formulations, the action is selected for maximising the expected value of the utility, i.e., for maximising the system performance. 17 sebastian thöns acta polytechnica ctu proceedings upa = max ak ( ∑ yk,m ∑ xl u(xl, ak, yk,m) · p (xl (ak, yk,m)) − c(ak) ) (14) for the notation of the action types, an index u for a utility action and an index s for a system state action are introduced. the objective function for a predicted utility action (pau) da may be written with the action implementation uncertainty p (yk,m, which is assigned to the utility. the action cost is here considered as deterministic for simplicity. upa = max ak ( ∑ yk,m ∑ yxl u(xl, ak) · p (yk,m) · p (xl) − c(ak) ) (15) the objective function for a predicted system state action (pau) da is written analogous with the action dependent system state probability p (xl(ak, yk,m)) and without the action influence on utility: variable mean cov characteristic value distribution reference load s with a reference period of 1 year 1.0 0.2 0.98 quantile gumbel e.g., jcss probabilistic model code [21], [22] representative for live and traffic loads load model uncertainty ms 1.0 0.1 logn. e.g., jcss probabilistic model code [21], part 3.9 (2001): model uncertainty for moments in frames, [22] resistance r 1.0 0.05 0.05 quantile logn. generic, jcss probabilistic model code [21], [22] resistance model uncertainty mr 1.0 0.05 logn. jcss probabilistic model code [21], part 3.9 (2001): model uncertainties, [22] deterioration d 0.1 0.02 logn. generic deterioration model uncertainty md 1.0 0.1 norm. generic damage resistance transfer function td td (md · d) = 1 − md · d generic safety factor γs 1.5 for one dominating action between 1.35 and 1.5, according to en 1990 (2010) safety factor γr 1.0 for steel between 1.0 and 1.5 according to material specific eurocodes consequence cf ∼ u (20.0, 40.0) [14] table 1. probabilistic system performance models. upa = max ak ( ∑ yk,m ∑ yxl u(xl) · p (xl(ak, yk,m)) · p (yk,m) − c(ak) ) (16) the system performance and objective functions for the introduced pa das are visualised with decision trees, generalised and summarised in figure 1. an action value can be quantified in analogy to an information value as the gain of an expected utility by a predicted action, vp a, i.e., as the difference between the maximised expected utility of a predicted action da (up a) and the system performance (usp ): vp a = up a − usp (17) the action value may be normalised (v p a) to the system performance without an action implementation, i.e.: v p a = (up a − usp ) / usp (18) 18 vol. 36/2022 structural assessment and expected utility gain action type and uncertainty space description probabilistic model cost and consequence model system state actions a a0 do nothing − − a1 strengthening of section 1.2 · γ 2.0% a2 replacement with higher strength 1.2 · e[rmat] 2.0% a3 repair md · d = 0 1.0% a4 load reduction s/1.2 0.5% a5 consequence reduction − 0.3% u: uniform distribution, tr: triangular distribution table 2. action type and cost models. action type and uncertainty space description probabilistic model implementation uncertainty y y1 strengthening performance p (y1) ∼ t r(0.95, 1.05, 1.10) y2 replacement performance − y3 repair performance p (y3) ∼ t r(0.9, 0.95, 1.0) y4 load reduction uncertainty p (y3) ∼ n (1.0, 0.1) y5 cons. reduction uncertainty p (y1) ∼ t r(0.85, 0.95, 1.05) n: normal distribution, tr: triangular distribution table 3. action uncertainty models. it should be noted that per definition, the action value analysis includes the cost of the action for consistency with the expected utility theorem. 4. exemplary study the predicted system state and utility action da is studied for an existing and deteriorated structure, for which the optimal risk reduction action is to be found. the system performance model includes the limit state function (2) representative for a highly correlated structural series system subjected to direct and indirect system failure consequences (section 4.1). the system state actions "section strengthening", "component replacement", "component repair" and "load reduction" as well as the utility action "consequence reduction" are probabilistically modelled and exemplified. 4.1. system performance model the system state model is represented with equation (2). for a semi-probabilistic safety concept, it can be shown that the section property can be decomposed to a central safety factor γ (see e.g., [23]), i.e., the product of the resistance and the loading safety factors (γr and γs , respectively), and the ratio rk of the characteristic resistance and loading values (rk and sk, respectively). the influence of the damage on the resistance is modelled with a damage transfer function td : z = γ · rk · td (md · d) = γr · γs · sk rmat, k · td (md · d) (19) the limit state function (2) becomes with focussing on cross section rather than material deterioration: x1 : gx1 = mr · γ · rk · td (md · d) · rmat − ms · s ≤ 0 (20) the probabilistic model and the references are summarised in table 1. 4.2. action model following the introduction of the action type and uncertainty model in section 2.2, a strengthening, a replacement, a repair, and a load reduction and a consequence reduction action are modelled (table 2). the section strengthening leads to an increase of 20% of the section property, whereas the replacement with a higher strength material leads to an increase of the expected resistance by 20% with a constant variability. both have a cost of 2.0%. the repair action sets the deterioration to 1.0 and costs 1.0%. the loading reduction action leads to decrease of the loading by 20% with the cost 0.5%. the consequence reduction sets the direct and indirect consequences in case failure to a uniform distribution between 10.0 and 20.0 with a cost of 0.3%. 19 sebastian thöns acta polytechnica ctu proceedings figure 2. system performance normalised relative action values without the action costs (left side, grey) and with action costs (left side, black) and total expected action cost and risks (right). the action implementation uncertainties are modelled continuous, i.e., with a triangular distribution accounting for a known range and most probable value (table 3). the strengthening action provides most probably a higher strengthening (1.05) than envisaged (table 2) within a range between 0.95 and 1.10. the component replacement is subjected to negligible action implementation uncertainties. the repair performance is subjected to on site execution uncertainties with a mode of 0.95 and a range of 0.9 to 1.0 of repairing the deterioration. the load reduction uncertainty is normal distributed and on average as expected and uncertain with a standard deviation of 0.1. the consequence reducing usage uncertainty is in the range of 0.85 to 1.05 with a mode of 0.95. 4.3. predicted action and action value analysis the results of the predicted action da and the action value analysis are depicted in figure 2. the load reduction action leads to the lowest expected costs and risks due to the comparably high load reduction including the uncertainties and the low action costs (figure 2, left). it also provides the highest action value including its costs, followed by the also relatively cheap consequence reduction action (figure 2, right, black). the load reduction action is identified as the optimal action in consistency with the expected utility theorem. the action values without their costs would be significantly higher for strengthening, replacement and repair (figure 2, right, grey). 5. information and integrity management value analysis the value of predicted information vp i and the value of information and actions are quantified as the gain of an expected and optimised utility by shi or shi and actions, respectively [1, 18]. the value of predicted information is then the difference between the expected and maximised utilities of a predicted information and action da (up ip a) and a predicted action da (up a): vp i = up ip a − up a (21) the difference between the expected and maximised utilities of a pipa da and the system performance (up a) is referred to as the value of the structural condition management encompassing information and actions vp ip a: vp ip a = up ip a − usp (22) the decision values may be normalised by the subtrahend in analogy to equation (18): v p ip a = (up ip a − usp ) / usp (23) it should also be noted that the information value can be calculated relative to system performance. in this case, the information value is divided by the expected system performance utility: 20 vol. 36/2022 structural assessment and expected utility gain v p i = (up ip a − up a) / usp (24) 5.1. predicted information and action da by threshold determination a general form of a pipa da can be derived for utility and system state actions (building upon the formulations in figure 1). for the shi acquirement strategies with the information outcomes zi,j and costs c(ii), the pipa da objective function in extensive form constitutes: up ip a = max ii ezi,j [ max ak eyk,m [ exl(ak ,yk,m |zi,j [u(xl, ak, yk,m] ] − c(ak) ] − c(ii) (25) in the following it will be shown that the above formulation (25) can be solved for continuous information outcomes by discretising to two complementary outcomes. continuous information outcomes are basically any measurement of a random variable contained in the system state function. per definition they are referred to direct shi . with consideration of the model uncertainty determination process in a multiplicative formulation [21], the model uncertainty can be understood as the measurable model outcome prediction normalised with the precise measurements of a structural reliability property. for a continuous model uncertainty outcome mm,i of information acquirement strategy with index i, equation (25) is readily rewritten: up ip a = max ii emm,i [ max ak eyk,m [ exl(ak ,yk,m |mm,j [u(xl, ak, yk,m] ] − c(ak) ] − c(ii) (26) the extensive form equation (26), can be rewritten with an obtained, i.e. posterior, information and predicted action decision analysis (oipa da) providing uoip a: up ip a = max ii emm,i [uoip a(mm,i)] − c(ii) (27) uoip a(mm,i) = max ak eyk,m [ exl(ak ,yk,m )|mm,i [u(xl, ak, yk,m) ] + c(ak) (28) the expectation operator in regard to the information can be written for continuous random variables: uoip a = max ii ∫ mm,i uoip a(mm,i) · fmm,i (mm,i) dmm,i − c(ii) (29) the measurement threshold tm,i,k is now found as the action-wise, i.e., ak dependent, expected posterior utility equality of "do nothing" and the predicted action, i.e.: tm,i,k : exl|tm,i,k [u(xl)] = eyk,m [exl(ak ,yk,m )|tm,i,k [u(xl, ak, yk,m]] + c(ak) (30) the integral in equation (29) can then be split up into two sub-integrals and the expected utility maximisation can be performed outside the integrals: up ip a = max ik ,ak   mm =tm,i,k∫ −∞ exl|mm [u(xl)] · fmm,i (mm,i) dmm,i + ∞∫ mm =tm,i,k [ eyk,m [ exl(ak ,yk,m )|mm [u(xl, ak, yk,m] ] + c(ak) ] · fmm,i (mm,i) dmm,i −c(ii)   (31) the integration can be rewritten for the model uncertainty density functions providing: up ip a = max ik ,ak   mm =tm,i,k∫ −∞ exl|mm [u(xl)] · dfmm,i (mm,i) + ∞∫ mm =tm,i,k [ eyk,m [ exl(ak ,yk,m )|mm [u(xl, ak, yk,m] ] + c(ak) ] · dfmm,i (mm,i) −c(ii)   (32) 21 sebastian thöns acta polytechnica ctu proceedings the two complementary outcomes zi,1 and zi,2 for action k = 1 are then defined with equations (33) and the integrals can be solved by calculation of the pre-posterior probability for direct information [12]: zi,1 : m ∣∣tm,i,1 −∞ p (zi,1) = fmm,i (tm,i,1) zi,2 : m ∣∣∞ tm,i,1 p (zi,2) = 1 − fmm,i (tm,i,1) (33) it should be noted that there is no further assumptions nor probabilistic models about information outcomes required [24]. 6. exemplary study the application of the threshold determination approach (equations (30), (32) and (33)) lead for the in section 4 introduced example to equation (34): given a realisation of the measurement, the threshold is to be found for which the optimal action changes to action implementation instead of do nothing. the measurement is subjected to a measurement uncertainty , which is normal distributed with a mean of 1.0 and a coefficient of variation of 0.03. the information cost equals 0.001, e.g. according to [11]. tm,k : p ( x1 ∣∣tm,k, mu) · e[cf ] = p(x1∣∣tm,k, mu , ak, yk,m) · e[cf (ak, yk,m)] + c(ak) (34) with the introduced limit state functions (see section 4.1), the threshold dependent probabilities of the failure state are calculated for damage information (equation (35)), loading information (equation (36)) and resistance information (equation (37)). p ( x1, tm,d,k, mu ) = p ( mr · z(tm,d,k · mu · d) · rmat − ms · s ≤ 0 ) (35) p ( x1, tm,r,k, mu ) = p ( mr(tm,r,k · mu ) · z(d) · rmat − ms · s ≤ 0 ) (36) p ( x1, tm,s,k, mu ) = p ( mr · z(d) · rmat − ms (tm,s,k · mu ) · s ≤ 0 ) (37) the solution of the first integral in equation ( 31), i.e. the calculation of the pre-posterior probability, is readily performed for damage information with the to the interval [0.0, tm,d,k] truncated (index t ) probability distribution of md [12]: p ( x1, tm,d,k, mu ) = p ( mr · z ( md [ 0.0, tm,d,k ]t · mu · d) · rmat − ms · s ≤ 0) (38) the full solution of equation (31) for damage, loading and resistance information is performed analogous. the pipa and value da in threshold formulation (equations (32) and (21) to (24)) result in the relative predicted action and information values for damage, load and resistance measurements (figure 3, left side). the damage measurement has no value for damage repair nor for replacement due to the nature of the actions. for the actions "load reduction" and "consequence reduction", damage monitoring leads regardless of and dependent on the damage realisation to higher expected utility and is thus not predictable. the integrity management strategies with the highest values (38.0% to 63.1%) are load and resistance measurements for the replacement, load reduction system state actions and the "consequence reduction" utility action. when considering separately the system performance related information values (figure 3, right side), it is observed that there are only positive information values for (1) load and resistance measurements in conjunction with the replacement action and (2) for load measurements for the strengthening and repair actions. only for the replacement action, measurement information contributes significantly to the condition management strategy value (see figure 3 and figure 2, right side). the optimal condition management strategies are thus replacement with load measurement (vp ip a = 63.1 % of the system performance), replacement with resistance measurement (vp ip a = 58.1 %of the system performance) and load reduction (vp a = 57.5 % of the system performance). 7. summary and concluding remarks this study builds upon structural condition assessment and provides a basis for the condition management by maximising the expected utility for information and actions before information acquirement and action implementation. the study is intended for progress towards consistent probabilistic, decision theoretical and built environment system relevant modelling of actions, information, expected utility quantification and value. 22 vol. 36/2022 structural assessment and expected utility gain figure 3. relative predicted action and information values for damage, load and resistance measurements. the study encompasses an introduction, probabilistic formulation, and exemplification (1) of system state and utility actions, (2) system state and utility action value analysis and (3) of a threshold formulation for predicted information and action decision analyses. the introduced probabilistic action modelling facilitates to distinguish between system state and utility actions, i.e., between actions influencing the reliability of the structural components and actions influencing the utility model. the action uncertainty modelling includes the distinction of an implemented or predicted action, the action implementation states and the action implementation uncertainties. with an action value decision analysis as introduced, the optimal action with consideration of the action costs can be identified before action implementation in consistency with the expected utility theorem. this is demonstrated for the system state actions strengthening, replacement, repair, load reduction and the utility action: consequence reduction. the threshold formulation for a predicted information and action decision analysis relies on the equality of the decision theoretical posterior action optimality condition, i.e., that the action is implemented according to the higher expected posterior utility. the threshold formulation is derived on the basis of direct shi and the measurability of model uncertainties. the identification of the threshold facilitates the quantification of the shi value for continuous measurement information and provides readily the decision rule, i.e., for which measurement outcomes the action implementation is optimal. the application of the information value threshold formulation and the action value analysis show exemplarily which structural integrity management strategies are optimal and how the action values and information values contribute to each strategy value. acknowledgements the kalix bridge (trv 2021/58089, ef 173793) and the condtionvalue (trv / bbt 2021/51600) projects funded by trafikverket, sweden, are gratefully acknowledged. references [1] h. raiffa and r. schlaifer. applied statistical decision theory. wiley classics library, originally published: boston : division of research, graduate school of business administration, harvard university, 1961. ed. new 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[24] s. thöns, a. a. irman, and m. p. limongelli, on uncertainty, decision values and innovation. the international conference on uncertainty in mechanical engineering (icume), darmstadt, germany, june 7 to 8, 2021. 24 https://doi.org/10.3390/infrastructures4030056 https://doi.org/10.1016/j.ress.2020.106814 https://doi.org/10.3389/fbuil.2020.00014 https://doi.org/10.1111/mice.12332 https://www.jcss-lc.org/publicationsjcss/ https://doi.org/10.1016/j.strusafe.2020.101957 https://doi.org/10.1080/15732479.2021.1890139 https://doi.org/10.1186/s43065-021-00027-0 https://doi.org/10.1016/j.marstruc.2020.102888 https://doi.org/10.1177/1475921720981833 https://doi.org/10.1016/j.strusafe.2018.09.002 acta polytechnica ctu proceedings 36:15–24, 2022 1 introduction 2 system performance and action modelling 2.1 system state analysis and system performance analysis 2.2 action modelling 3 predicted action decision analysis and action value analysis 4 exemplary study 4.1 system performance model 4.2 action model 4.3 predicted action and action value analysis 5 information and integrity management value analysis 5.1 predicted information and action da by threshold determination 6 exemplary study 7 summary and concluding remarks acknowledgements references my title https://doi.org/10.14311/app.2022.36.0253 acta polytechnica ctu proceedings 36:253–260, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague simplified modelling of the performance of concrete tunnels during fire and post-fire damage classification ruben van coilea, ∗, balša jovanovića, ranjit kumar chaudharya, xavier deckersb, andrea lucherinia a ghent university, department of structural engineering and building materials, technologiepark-zwijnaarde 60, 9052 gent, belgium b fire engineered solutions ghent a jensen hughes company, oudenaardsesteenweg 32 g, 9000 ghent, belgium ∗ corresponding author: ruben.vancoile@ugent.be abstract. the performance of concrete tunnel structures during and after fire is not well understood. this is an obstacle to the adoption of risk-based approaches for fire safety design of tunnel structures. upon the request of the belgian fire safety consultancy fesg, a simplified assessment of the collapse probability and post-fire damages for a reference tunnel structure has been made. the structural system is modelled through 2d beam finite elements, where spalling rates have been assumed based on available literature data. structural stability is verified for both the heating and cooling phases of the fire. in those cases where the structure survives up to burnout, the residual deformations and thermal damage to the tunnel structure are assessed. keywords: damage, failure probability, fire, spalling, tunnel. 1. introduction the structural fire design of concrete road tunnels is commonly done with reference to prescriptive guidance and regulations. the actual performance achieved in case of a fire is not well understood, both in terms of the probability of failure (collapse during fire), as well as the probability density function of the damage post-fire (permanent damage and residual deflections, repair costs). this hampers the adoption of risk-based approaches to the fire safety design of road tunnels, and is an obstacle for risk-based codecalibration. the belgian fire engineering consultancy fesg has been developing a simplified risk assessment tool for tunnel structures. this risk assessment requires an evaluation of the probability of structural failure during fire and of the extent of the post-fire damage and downtime. this will allow for the comparison of design alternatives, e.g. with respect to active and passive fire safety measures. for further details on the project and a conceptual framework for decision making for tunnel fire safety, see [1]. in cooperation with fesg, a simplified assessment of the performance of a reference concrete tunnel structure during and after fire has been made. a 2d beam finite element model is adopted, taking into account the effect of concrete spalling through nominal spalling rates. the fire is modelled through a heating phase according to the nominal rws curve, followed by a decay phase based on limited experimental data. details of the approach and key outcomes are provided in the following. 2. modelling approach 2.1. reference tunnel section the reference tunnel cross section is visualized in figure 1, with tunnel lining cross-sectional characteristics (per unit length) listed in table 1. the cross-section is inspired by an existing design consisting of two one-directional tubes and a central connecting tube for evacuation purposes and maintenance. for simplicity (calculation time), all walls are identical and no reinforcement curtailment has been modelled. soil resistance to outwards movement of the tunnel is modelled through springs. a structural mesh of 0.5 m is applied. a fine thermal mesh is applied (5 to 10 mm element width perpendicular to the heat transfer direction). the structural modelling is done using safir [2]. other input parameters for the analysis (material and mechanical properties of concrete and reinforcement) are listed in table 2. the adopted concrete model is an extension to the en 1992-1-2:2004 [3] model, taking into account explicit modelling of transient creep [4]. concrete tensile strength is neglected. the concrete compressive strength and reinforcement yield strength are considered in the model with their expected values. this is the preferred approach for implementation in the larger risk-based framework. furthermore, due to the complex structural behaviour of the tunnel system, it is not a priori clear if a larger or smaller value of the material strength and stiffness is "more conservative". the concrete poisson coefficient is an artefact from model development. a lower temperature-dependent poisson coefficient can 253 https://doi.org/10.14311/app.2022.36.0253 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en r. v. coile, b. jovanović, r. k. chaudhary et al. acta polytechnica ctu proceedings figure 1. section of the reference tunnel and dimensions (lining axis positions). the blue lines are springs modelling the soil interaction. parameter ceiling and floor walls thickness [mm] 600 410 concrete cover [mm] 70 100 reinforcement axis distance to concrete surface [mm] 75 105 reinforcement area (single layer) [mm2] 5300 �25 − 100 2010 �16 − 100 table 1. tunnel cross-sectional characteristics (per unit length). be recommended, in accordance with [5]. parameter value concrete compressive strength [mpa] 42 concrete tensile strength [mpa] 0 concrete poisson coefficient [−] 0.3 reinforcement yield strength [mpa] 560 reinforcement initial modulus of elasticity [gpa] 210 reinforcement poisson coefficient [−] 0.3 table 2. material parameters. upon cooling, concrete loses an additional 10% of strength in accordance with en 1994-1-2:2005 recommendations. the reinforcement fully recovers upon cooling. the latter is a simplification that overestimates reversibility for reinforcement which has exceeded 600◦c [6]. this simplification is in line with the state-of-the-art, but means that cooling results for reinforcement that has exceeded 600◦c need to be interpreted with care. fundamental research is needed to develop a model for the material behaviour during the cooling phase. the tunnel is loaded with 1 m of soil (weight 17.5 kn/m3) and 2 kn/m2 permanent load at ground level. the lateral earth pressure coefficient is set at 0.4. the tunnel is considered to be above the ground water table (no upwards water pressure). these values were chosen to represent a city tunnel for through traffic. a characteristic imposed load of 30 kn/m2 is considered. the mean imposed load is assessed as 0.2 times the characteristic value, in analogy with building design [7]. improved live load modelling for geotechnical structures, and the consideration of the local geotechnical parameters are recommended. 2.2. fire exposure as the tunnel is symmetrical, fire is considered in one of the road tubes only (see figure 1). the likelihood independent simultaneous occurrence of fire in both tubes is considered negligible. within the larger risk assessment tool, a detailed assessment of fire severity is made, taking into account fire spread and the interaction with active fire safety measures. for the structural performance evaluation, these fire severities are mapped onto an equivalent heating duration for the nominal rws fire curve [8]. the following nominal durations are considered: 15/30/60/120/180 min. the fire exposure is applied to the walls and ceiling, considering a convective heat transfer coefficient of 50 w/(m2k) and emissivity of 0.7, based on en 19921-2:2004 [3]. the decay phase is modelled through equation 1, with θc, θref and θmax respectively the fire curve temperature in the cooling phase, the reference ambient temperature (20◦c) and the maximum rws fire temperature during the heating phase. the time in minutes since the nominal start of the fire and the time of heating phase duration are denoted by t and tmax, while the speed of the decay is governed by the decay parameter b. to evaluate the decay parameter b in equation 1, the decay phases of 19 test measurements were extracted from [9]. an average value of 0.025 min−1 was estimated and adopted (mean value preferred for cost-benefit evaluations in a later stage). for concrete structures, the cooling phase behaviour is of great importance, both due to the delayed heating of the cross-section, as well as 254 vol. 36/2022 modelling of concrete tunnels damage by fire figure 2. example fire curves. rws heating phase and exponential decay phase. temperatures to be interpreted as adiabatic surface temperatures. due to the fire-induced permanent deformations resulting in a possible force redistribution. the obtained temperature-time curves are visualized in figure 2. θc − θref = (θmax − θref ) e−b(t−tmax) (1) 2.3. concrete spalling spalling is modelled considering nominal spalling rates, as recommended in [10] and applied by [11]. specifically, the start of spalling is fixed at 1 min after the nominal start of the exposure, followed by a constant spalling rate of, for example, 3 mm per minute. spalling is stopped as soon as either (i) the reinforcement layer is reached, or (ii) when the fire enters the decay phase. the spalling is implemented in safir through an automated procedure, as described in [11]. considering the current state-of-knowledge, it is recommendable to consider all unprotected concrete to be prone to spalling. earlier definitions of spalling-free concrete mixes have later been shown susceptible to spalling after all, see e.g. [12]. especially in the case of severe fire exposures as adopted for tunnels, (limited) spalling of unprotected concrete can be assumed. the addition of polypropylene fibres is considered to reduce the concrete propensity to spall [13]. most literature sources ascribe this to an increased porosity (and thus the reduction in pore pressures) during fire. concrete with polypropylene fibres is assigned a spalling probability of 40%. this number is based on a study of spalling occurrence in an unpublished test campaign and should be considered as engineering judgement. for protected concrete, spalling is considered not to occur. this does not imply that local spalling is deemed impossible in case of protected concrete. the occurrence of spalling across the entire exposed surface is however deemed very unlikely. hua et al. [11] report spalling rates for tunnels of up to 5 mm/min. in the current study nominal spalling rates of 1/2/3/3.75/5 mm/min are considered, as well as a no-spalling case. the spalling severity figure 3. spalling rate data listed by hua et al. (2021) and fitted gamma distribution. spalling rate [mm/min] probability 1.00 0.091 2.00 0.313 3.00 0.292 3.75 0.139 5.00 0.165 table 3. assessment of conditional probabilities for the spalling severity. probabilities (conditional probability given spalling) are derived from the data listed by [11]. a theoretical distribution describing the spalling rate probability density is fitted on the data whereby each literature data point is treated equally. applying the akaike information criterion (aic) [14], a gamma distribution (equation 2, with γ the gamma function, and x the spalling rate) with shape factor k = 3.6 and scale factor θ =0.7 is chosen as the best fit (see obtained fit in figure 3). based on this distribution, probabilities for each spalling rate used in this study are calculated and presented in table 3. these probabilities have been calculated considering the nominal spalling rates as upper bounds for their respective interval (with the exception of the highest nominal spalling rate of 5 mm/min). f (x) = 1 γ(k) · θk xk−1e(x/θ) (2) 2.4. post-fire damage assessment reinforced concrete structures exhibit permanent damage following fire [6]. concrete exhibits a permanent loss of strength and stiffness following heating to elevated temperatures and subsequent cooling, and also reinforcement exhibits a permanent loss of 255 r. v. coile, b. jovanović, r. k. chaudhary et al. acta polytechnica ctu proceedings damage states thermal damage repairability ds0 d300 ∼= 0 no repair required ds1 0 < d300 < c/10 clean surface and replace damaged concrete ds2 c/10 ≤ d300 < c clean surface and replace damaged concrete ds3 c ≤ d300 < d/4 clean surface and replace damaged concrete ds4 d/4 ≤ d300 < d/2 demolish and reconstruct table 4. thermal damage states, with c the concrete cover and d the thickness of the cross-section. damage states structural damage ceiling repairability structural damage wall repairability ds0 ∆ s/l < 1/240 no repair required ∆ c/h < 0.5% no repair required ds1 1/240 ≤ ∆ s/l < 1/120 no repair required ∆ c/h ≥ 0.5% demolish andreconstruct ds2 1/120 ≤ ∆ s/l < 1/60 demolish and reconstruct −ds3 1/60 ≤ ∆ s/l demolish and reconstruct table 5. structural damage states, with ∆ s the mid-span residual vertical displacement, l the ceiling span, ∆ c the mid-height residual horizontal displacement, and h half of the wall height. figure 4. axial force (top; compression blue) and bending moment (bottom) before the fire. strength following exposure to temperatures in excess of 600◦c [6]. both these damage effects are on a material level. the fire exposure also results in fire-induced forces (restraining effects), permanent thermal strains, and load redistribution due to changes in stiffness. these effects result in residual deformations post-fire. these residual deformations may jeopardize the serviceability of the tunnel, and also result in second order load effects. these latter damage effects are on a system (structural) level. the concept to pre-evaluate the repair costs of concrete structures following fire exposure is relatively new. a recent in-depth investigation for concrete structures has been presented in [15]. the evaluation of concrete fire damage through the 300◦c isotherm (d300) and residual deformations is recommended. these damage parameters are then used to determine the damage class and repair feasibility. the study by ni and gernay also elaborates an approach for repair cost of fire-damaged structures based on the damage states. specifically, rsmeans data is used to inform the cost evaluations. a similar approach as proposed by ni and gernay is adopted in the following to evaluate the damage class and repair cost for fire-damaged tunnel structures. the same limit states as proposed by [15] are used here to determine the damage of the structure. table 4 and table 5 below gives the damage states and repairability limit states for thermal and mechanical damages for concrete structures following a fire event. 256 vol. 36/2022 modelling of concrete tunnels damage by fire figure 5. axial force (top; compressive blue) and bending moment (bottom) after approx.. 60 min of fire exposure (beginning of the cooling phase). figure 6. axial force (top; compression blue) and bending moment (bottom) after 12 hours of fire exposure (end of the fire event simulation. for the damage states of the tunnel ceiling, the damage classification for slabs is adopted, while column limits are adopted for the walls. this is based on the consideration that the tunnel cross-section can be viewed as a 2d frame. 3. results 3.1. example case analysis 60 min rws heating phase, no spalling in this case study, the walls and ceiling in the left tube are exposed to fire from the interior only. these linings can be considered thermally thick as the cooling conditions on the unexposed side (e.g., adiabatic, or cooling to ambient) do not influence the results. upon the start of the cooling phase, the hot exposed surface quickly starts to cool as well. the heat wave however 257 r. v. coile, b. jovanović, r. k. chaudhary et al. acta polytechnica ctu proceedings figure 7. thermal damage state (ceiling and wall) in function of rws heating phase duration and nominal rate of concrete spalling. figure 8. structural damage state (ceiling and wall) in function of rws heating phase duration and nominal rate of concrete spalling. continues to penetrate into the interior of the lining. at 130 minutes the ceiling bottom reinforcement reaches its maximum temperature of approximately 270◦c. this maximum temperature is within the range where the reinforcement can be considered to fully recover [6]. similar results are obtained for the fire-exposed walls. as the concrete cover is larger for the walls, the maximum reinforcement temperature at the exposed side is lower (189◦c, attained at approximately 168 minutes). the depth of the 300◦c isotherm reaches a maximum around 113 minutes with a depth of 60-65 mm. this result applies to both the ceiling and wall. the thermal damage classification is ds2 both for the ceiling and wall. the tunnel section withstands the modelled exposure without structural failure. however, the fire exposure does result in permanent deformations and load redistributions. importantly, to obtain a full view on the structural response, the calculation needs to be continued for beyond the end of the heating phase. in figure 4, the axial force (top) and bending moment (bottom) in the tunnel section are visualized at the start of the exposure (compression blue, tension yellow colour; bending moment drawn at the tension side). figure 5 visualizes the same at approximately 60 min, and figure 6 at 12 hours since the start of the nominal fire exposure. in these figures, displacements are visualized with a factor 10. the maximum ceiling mid-span deformation is obtained at approximately 115 min since the start of the exposure (approximately 55 mm). the recovery of the deformation takes multiple hours. the residual total mid-span deformation is 36.4 mm. the mechanical damage classification for the ceiling is ds0. the same applies to the wall. 3.2. damage classification summary none of the simulated tunnels fail during the fire exposure. the possibility of failure has however been confirmed through simulations with characteristic loading. in the current evaluation, expected values for the load conditions have been considered as indicated in section 2.1. figure 7 and figure 8 show the matrix plots for the thermal and mechanical damage for different rws fire duration and nominal spalling rates. the plots also include damage classifications for fire protected structures (i.e. tunnels cladded with passive fire protection boards with a nominal rating of 1h or 2h protection). combining the results of figure 7 and figure 8 with the spalling probabilities listed in section 2.3 above, damage class probabilities are obtained in function of the equivalent fire severity (figure 9 for the ceiling damage classes for unprotected tunnel linings). figure 9 highlights how also limited fire exposure results in a damage class of at least ds2, and a high proba258 vol. 36/2022 modelling of concrete tunnels damage by fire figure 9. structural damage state probabilities (probability of damage class dsx or higher) for unprotected tunnel ceiling in function of rws heating phase duration. note that the damage class is evaluated for the heating durations 15/30/60/120/180 minutes only. bility of class ds3. probabilities of ds4 are however observed only from 120 minutes heating duration onwards (note that only a limited number of discrete heating durations are modelled here). the results of figure 9 are subsequently combined with the rws heating phase duration probabilities provided by the fesg fire severity model, thus resulting in damage class probabilities for a given tunnel with specific traffic intensity and active/passive fire safety measures. within the larger risk assessment tool, these damage class probabilities are subsequently translated into assessments of the downtime and repair costs. 4. conclusions the structural fire performance of a reference concrete tunnel has been evaluated, taking into account different rws heating durations and nominal spalling rates. the total simulated time is 12 hours to explicitly evaluate the cooling of the structure. the importance of an explicit evaluation of the cooling phase behaviour has been demonstrated through an example, and the probabilities of achieving set thermal and structural damage states post-fire have been assessed. the results will be incorporated in a simplified risk assessment tool, allowing to compare the risk profile of design alternatives for concrete road tunnels. acknowledgements this study has been made possible through funding from fesg, supported by vlaio (flanders innovation & entrepreneurship). 5. bibliography [1] r. van coile, m. schepers, m. pachera, m., et al. risk acceptance framework for specifying fire safety measures in tunnels. 13th international conference on structural safety and reliability (icossar 2021), 20-24/06 shanghai, china, 2022. [2] j.-m. franssen, t. gernay. modeling structures in fire with safir®: theoretical background and capabilities. journal of structural fire engineering 8(3):300-23, 2017. https://doi.org/10.1108/jsfe-07-2016-0010. [3] en 1992-1-2:2004. eurocode 2: design of concrete structures part 1-2: general rules structural fire design. european standard, 2004. https://www.phd.eng.br/wp-content/uploads/2015/ 12/en.1992.1.2.2004.pdf. [4] t. gernay, j. m. franssen. a formulation of the eurocode 2 concrete model at elevated temperature that includes an explicit term for transient creep. fire safety journal 51:1-9, 2012. https://doi.org/10.1016/j.firesaf.2012.02.001. [5] t. gernay, a. millard, j.-m. franssen. a multiaxial constitutive model for concrete in the fire situation: theoretical formulation. international journal of solids and structures 50(22-23):3659-73, 2013. https://doi.org/10.1016/j.ijsolstr.2013.07.013. [6] fib. fire design of concrete structures structural behaviour and assessment. fib bulletin 46. the international federation for structural concrete. lausanne, switzerland, 2008. [7] b. jovanović, r. van coile, d. hopkin, et al. review of current practice in probabilistic structural fire engineering: permanent and live load modelling. fire technology 57(1):1-30, 2020. https://doi.org/10.1007/s10694-020-01005-w. [8] a. j. breunese, c. both, g. m. wolsink, g.m. fire testing procedure for concrete tunnel linings. tno-rapport 2008-efectis-r0695, bleiswijk, the netherlands, 2008. [9] h. ingason, s. gustavsson, m. dahlberg. heat release rate measurements in tunnel fires. brandforsk project 723-924 (sp report 1994:08). swedish national testing and research institute, boras, sweden, 1994. [10] fib. performance-based design of concrete structures (committee draft 2021-06-24). the international federation for structural concrete. lausanne, switzerland, 202. [11] n. hua, a. tessari, n. elhami khorasani. characterizing damage to a concrete liner during a tunnel fire. tunnelling and underground space technology 109, 2021. https://doi.org/10.1016/j.tust.2020.103761. [12] t.g. van der waart van gulik, a. j. breunese, r. jansson, r., et al. spalling behaviour of a non-spalling qualified concrete. 4th international workshop on concrete spalling due to fire exposure, 08-09/10 leipzig, germany, 2015. [13] c. maluk, l. bisby, g. p. terrasi. effects of polypropylene fibre type and dose on the propensity for heat-induced concrete spalling. engineering structures 141:584-95, 2017. https: //doi.org/10.1016/j.engstruct.2017.03.058. [14] h. akaike. a new look at the statistical model identification. ieee transactions on automatic control 19(6):716-23, 1974. https://doi.org/10.1109/tac.1974.1100705. 259 https://doi.org/10.1108/jsfe-07-2016-0010 https://www.phd.eng.br/wp-content/uploads/2015/12/en.1992.1.2.2004.pdf https://www.phd.eng.br/wp-content/uploads/2015/12/en.1992.1.2.2004.pdf https://doi.org/10.1016/j.firesaf.2012.02.001 https://doi.org/10.1016/j.ijsolstr.2013.07.013 https://doi.org/10.1007/s10694-020-01005-w https://doi.org/10.1016/j.tust.2020.103761 https://doi.org/10.1016/j.engstruct.2017.03.058 https://doi.org/10.1016/j.engstruct.2017.03.058 https://doi.org/10.1109/tac.1974.1100705 r. v. coile, b. jovanović, r. k. chaudhary et al. acta polytechnica ctu proceedings [15] s. ni, t. gernay. a framework for probabilistic fire loss estimation in concrete building structures. structural safety 88, 2021. https://doi.org/10.1016/j.strusafe.2020.102029. 260 https://doi.org/10.1016/j.strusafe.2020.102029 acta polytechnica ctu proceedings 36:253–260, 2022 1 introduction 2 modelling approach 2.1 reference tunnel section 2.2 fire exposure 2.3 concrete spalling 2.4 post-fire damage assessment 3 results 3.1 example case analysis 60 min rws heating phase, no spalling 3.2 damage classification summary 4 conclusions acknowledgements 5 bibliography bibliography acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0015 acta polytechnica ctu proceedings 40:15–21, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague measurement and evaluation of cement paste porosity by electrochemical impedance spectroscopy jana chalupová a, ∗, jiří němeček 1,a, vojtěch hybášekb, vojtěch pommerc, jiří němeček 2,a a czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic b university of chemistry and technology, department of metals and corrosion engineering, technická 5, 166 28 prague 6, czech republic c czech technical university in prague, faculty of civil engineering, department of material engineering and chemistry, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: jana.chalupova@fsv.cvut.cz abstract. porosity is an important part of cement paste microstructure, which significantly influences the mechanical properties (especially durability and strength) of cement-based materials. electrochemical impedance spectroscopy (eis) is a non-destructive method used to measure the pore system utilizing a range of frequencies of electric current, which is fed to the cement paste by stainless steel electrodes. the pore volume is obtained from the pore resistance and the matrix capacitance measured by eis. this paper deals with the evaluation of porosity based on resistance values from eis on a sample of pure cem i 42.5r portland cement paste at 3 different hydration times (age 7, 14, 28 days). keywords: electrochemical impedance spectroscopy, cement paste, porosity. 1. introduction cement is one of the most consumed materials in the world. therefore, many scientists are working on researching, developing, and improving cement-based materials. porosity is one of the observed properties at the micro-scale that influences strength and durability [1]. generally, three types of pores can be found in hydrated cement paste – capillary pores, gel pores, and interlayer space. the pore characteristics differ in size, shape, and distribution in the solid matrix. porosity measurement depends on all these characteristics [2]. the porosity of cement-based material can be accessed either by direct or indirect methods. direct methods are able to measure the pore sizes from the largest pores in the order of a few millimeters to hundreds or tens of nanometers. direct methods analyze pore structure from 2d or 3d images of the sample. therefore, these methods depend on the resolution of an image (pixel size). these methods include 2d images of scanning electron microscopy (sem), 3d images of x-ray microtomography or x-ray nanotomography, and laser scanning confocal microscopy (lscm) [3]. indirect methods use liquid (such as water, gas, mercury, helium, etc.) inside cement paste as a probe. 1ph.d. at czech technical university in prague, orcid: 0000-0002-5635-695x. 2professor at czech technical university in prague, orcid: 0000-0002-3565-8182. the size of the measurable pores is limited by the lower (tens of nanometers) and upper boundary (hundred of micrometers). indirect methods of measuring open porosity include mercury intrusion porosimetry (mip), water saturation, and pycnometry [3–5]. electrochemical impedance spectroscopy (eis) can be ranked as the indirect method because a liquid conduct an electric current and thus is involved in the measurement. using eis not only detects the porosity and microstructure of concrete [6–9], but also the hydration and shrinkage process [10], corrosion process of reinforced concrete [11, 12], the diffusion of chloride in concrete [13, 14] and the influence of mineral admixtures on cement-based material [15]. despite some efforts of porosity characterization by eis [6–9] the methodology is still limited by many uncertainties. thus, this paper deals with the eis measurement on pure cement paste samples, evaluation of porosity, and comparison with porosity determined by helium pycnometry. 2. electrochemical impedance spectroscopy 2.1. equivalent circuit models for concrete electrochemical impedance spectroscopy (eis) or otherwise known as alternating current impedance spectroscopy (acis), is a non-destructive method used to measure the resistance of materials, including cement 15 https://doi.org/10.14311/app.2023.40.0015 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en j. chalupová, j. němeček, v. hybášek et al. acta polytechnica ctu proceedings paste and concrete. electric current (usually alternating current ac) is applied to the samples by a pair of electrodes embedded in the material or attached to their opposite surfaces. the transmitted current with the periodic alternating signal can be decomposed into a series of harmonic signals of sinusoidal shape with different frequencies due to fast fourier transform [16, 17]. the measurement is taken from the highest frequencies to the lowest, and the impedance is recorded [6, 12]. cement paste cannot be considered as a single electrical resistor in eis measurements because the material consists of a bulk matrix and pores. pores have different shapes, sizes, and distributions in the cement paste structure. for example, capillary pores are connected and with an oblong shape in contrast to a spherical enclosed pores [2]. electrochemical impedance spectroscopy allows measuring only the capillary pores, and three conductive paths are distinguished as shown in the figure 1 [6]: • continuous conductive path (ccp) – a series of capillary pores connected by pore necks form ccp. the impedance of this path z ccp equals the resistance of interconnected capillary pores rccp: z ccp = rccp. (1) • discontinuous conductive path (dcp) – the continuity of a capillary path is disturbed by a discontinuous point (dp) formed by a thin layer of cement paste. dp is considered to be a double parallel capacitor with electric capacitance c dp. the resistance of the capillary path is labeled as rcp. the impedance of the path z dcp is composed of c dp and rcp connected in series as shown the following equation: z dcp = rcp + c dp. (2) • insulator conductive path (icp) – the pores and voids do not form the largest part of cement paste. it is made up of a bulk matrix that acts as an electric insulator. the matrix becomes charged as a result of the current passing through the sample. therefore it is considered a double parallel capacitor with capacitance c mat. although cement paste is not a perfect insulator, its resistance rmat can be neglected if the sample is not frozen or dried. thus, the impedance of this path z icp is just equal to the c mat: z icp = c mat. (3) according to the typical nyquist diagram (plot of real versus imaginary part of impedance), capacitive loops occur in cement-based materials as shown in figure 2b. these loops can be replaced by a parallel series of resistors (r) and capacitors (c ). an equivalent electric circle containing r and c is necessary to use in order to evaluate the results of eis measurement [18, 19]. many equivalent circuits have been published and are summarized in [19]. discontinuous conductive path (dcp) continuous conductive path (ccp) insulator path (icp) discontinuous point (dp) pore solution cement paste conductive path figure 1. simplified microstructure of cement paste with illustrated conductive paths. in this paper, two different models of equivalent electric circuits published by guangling song in [6] were used. the first option is the equivalent circuit model (ec), which is composed of all above mentioned paths, as shown in figure 2a. the capacitance of the solid matrix of the cement paste is relatively low. therefore the second option assumes the simplification that c mat does not have to be considered in the second equivalent model, as shown in figure 3. furthermore, the electric circuit can be simplified into the simplified equivalent circuit model (sec) as shown in figure 2c. comparing the equivalent models in figures 2c and 3, the following relations can be deduced [6]: r0 = rcp · rccp/(rcp + rccp), (4) r1 = rccp2/(rcp + rccp), (5) c1 = (1 + rcp/rccp)2 · c dp, (6) where r1 is resistance of (continuous and discontinuous) pores, c1 is a capacitance of bulk cement (including c dp), r0 is an offset resistance from the origin on real axis of nyquist diagram. the theoretical nyquist spectrum of the eis consists of arcs. the most important arc of the sec model is the arc of diameter r1, see figure 2d. the position of the center of the arc on the real axis z is ensured by rotation by the depression angle (α) [20], which is related to the pore size distribution and the others imperfection of the sample [12]. r0 can be neglected because of high-frequency measurement (nearly to the origin on the real axis of the nyquist plot) is negatively influenced by surrounding phenomena and the accuracy limitation of most electrochemical equipment [6, 20, 21]. 16 vol. 40/2023 measurement and evaluation of cement paste porosity . . . rcp cdp cmat rccp bulk matrix connected pores disconnected pores zicp zdcp zccp re z [ω] -i m z [ω ] b) nyquist spectrum based on ec impedance r0 r1 a) equivalent circuit (ec) r0 c1 r1 [ω] d) nyquist spectrum based on sec impedance r0 r1 c) simplified equivalent circuit (sec) re z -i m z [ω ] figure 2. equivalent circuit model (ec), theoretical nyquist eis spectrum based on a equivalent model ec, simplified equivalent circuit model (sec) and theoretical nyquist eis spectrum based on a equivalent model sec [6]. rcp cdp rccp connected pores disconnected poreszdcp zccp figure 3. simplified equivalent circuit model for cement paste used to measure of eis [6]. 2.2. porosity calculation from eis measurements archie’s law is the most widely used relationship between the resistivity and porosity of porous material. the relationship was originally derived on the sandstone filled with brine [22] and is defined by the equation: f = a · ϕ−m0 , (7) where f is formation factor, a is a coefficient, ϕ0 is the capillary porosity, m is archie’s index. in cement-based materials, the resistivity is dependent on sample dimensions and electrode positions. thus, it is more convenient to convert the resistance to effective electrical conductivity σeff as a function of electrode size and position: σeff = l/(r · s), (8) where l is the distance between the electrodes in the direction of current, s is the cross-sectional area of the electrode embedded in the cement paste. σeff = c · σ0 · ϕ0m , (9) where σ0 is the conductivity of conducting medium, c is a constant depending on the saturation of the sample (assumed to be 1.0 for fully saturated samples), ϕ0 is the pore volume fraction, m is archie’s index. the exponent m reflects pore complexity and tortuosity and has been found in the range 1.5–4.0 [13, 18]. 3. materials and methods 3.1. sample preparation cement paste samples were made from cem i 42.5r with a water to cement ratio of 0.4. a fresh mixture was poured into a silicone-lubricated formwork with dimensions of 54×30×11.7 mm (length, width, height). additionally, a pair of electrodes from a stainless steel plate was put longitudinally into the formwork. electrodes were inserted 3 mm above the formwork bottom and 1 mm from formwork edges. the distance between electrodes was set to 10 mm. the scheme of sample setup for eis measurement is shown in figure 4. moreover, the same samples without electrodes were prepared for porosity measurement. the samples were demoulded after 24 hours and placed in 0.5 % limewater solution where retained until the measurement. 17 j. chalupová, j. němeček, v. hybášek et al. acta polytechnica ctu proceedings 11 .7 15 9. 7 3 r e f e r e n c e e l . w o r k in g e l . c o u n t r e l . w o r k in g e l . 52 54 1 1 10 10 10 30 [mm] embedded electrode cement paste figure 4. an illustration of the sample with embedded and connected electrodes. 3.2. eis measurement eis measurement was performed using zahner zennium x device and thalesxt usb software with a frequency range of 12 mhz–100 hz and 10 steps per decade. the amplitude of sinusoidal voltage was set to 10 mv. the use of a larger potential amplitude is not recommended due to possible changes in the surfaces of the samples [17]. the cables connected two working electrodes to one of the embedded electrodes. the reference electrode and the counter electrode were connected to the other of the embedded electrode [23]. short coaxial cables were used to reduce the influence of surrounding phenomena and the measurement noise [17]. to ensure consistent measurement results, it was necessary to prevent the drying of the saturated samples. thus, samples were partly submerged in the water while simultaneously avoiding contact between electrodes and water. otherwise, the electric current entering the samples would pass through the water, and the results would not correspond to the cement paste. the samples were measured at 3 different ages of hydration – 7, 14, and 28 days. the eis measurement was taken at five-minute intervals until the values stabilized. 3.3. porosity measurement the porosity of the cement paste was determined by helium pycnometry and mercury intrusion porosimetry. after a saturation period of 7, 14, and 28 days, the samples were cut into 2 mm thick slices and dried at 50 °c. subsequently, sample density, ρ was measured by helium pycnometer thermo scientific atc evo. the bulk density, ρbulk of the samples was determined by the gravimetric method. further, the porosity of the sample was calculated from the equation: ϕ0,p = (1 − (ρbulk/ρ)) · 100. (10) 4. results and discussion the example of the eis results for the c-14d and c-28d samples is shown in figure 5 as the nyquist 150 200 250 300 0 50 100 150 200 re z [ω] -i m z [ω ] a) c-14d exp. data cem. paste exp. data electrode exp. data noise ec fit sec fit 150 200 250 300 0 50 100 150 200 re z [ω] -i m z [ω ] b) c-28d exp. data cem. paste exp. data electrode exp. data noise ec fit sec fit figure 5. experimental nyquist eis spectrum with calculated ec and sec model fits for samples c-14d and c-28d. note that only data corresponding to cement paste (blue marks) were used for fitting. spectrum. before the resistance and capacitance evaluation, experimental data were corrected and cut off at both ends. that is because the values at low frequencies correspond to the electrodes’ resistance and their contact with the material, and values at high frequencies are inaccurate due to the limitations of the measuring device. further, the data were fitted by the simplex algorithm with the application of ec and sec models [6]. the samples were remeasured over time until the resistance values were stabilized. the stabilization period took between the 15–30 minutes, as illustrated in figure 6. table 1 summarizes the resistances and capacities corresponding to the fits based on the ec and sec models, as well as degree rotation α. the results clearly show almost no differences between resistances rccp (ec) and r1 (sec). therefore, both values correspond to the continuously connected pores, indicating that c mat for the pure cement paste 7–28 days old can be neglected as considered by the simplified equivalent model. also, both rccp and r1 are increasing with samples age. such a behavior corresponds 18 vol. 40/2023 measurement and evaluation of cement paste porosity . . . sample equivalent circuit – ec simplified equivalent circuit – sec rccp rdcp rdp cmat r0 r1 c1 α [ω] [ω] [pf] [pf] [ω] [ω] [pf] [rad] c-7d 201.3 1290 29.1 38.7 0 204.1 41.9 0.81 c-14d 215.2 1577 29.2 40.7 0 217.3 42.0 0.82 c-28d 274.5 1875 21.7 41.6 0 276.8 45.5 0.86 table 1. resistance and capacitance of cement paste samples evaluated with different equivalent circuit models: ec, sec. 0 5 10 15 20 25 30 200 225 250 275 300 325 time [min] r c c p [ω ] c-14d c-28d figure 6. a stabilization period of cement paste resistance at different hydration times. the ec model evaluated rccp. well to the decreasing porosity due to the ongoing hydration reaction, during which the newly created hydration products occupy space originally formed by capillary pores. thus, creating more discontinuous paths results in increasing rdcp over time. the density results measured by helium pycnometry and bulk density are summarized in table 2 along with the values of calculated porosity (ϕ0). again the decrease in porosity corresponds to the hydration process. initially, the porosity from the eis measurement could not be directly calculated using equation (9) and porosity obtained from helium pycnomtery was used for calibration. another expression, the effective electrical conductivity was easily calculated from resistances rccp and r1. the conductivity of the pore solution σ0 was calculated by the analytical relationship of snyder [24], which is dependent on the w/c ratio, degree of hydration (doh), and the amount of na+, k+, oh− concentration. the doh was estimated by cemhyd3d model [25]. thus, the only unknown term remains archie’s index m. it was found that parameter m is almost identical for both equivalent circuit models ec and sec and also does not vary with hydration time ranging from 7 to 28 samples age. therefore, it is possible to calculate the average value sample ρ ρbulk ϕ0,p [kg/m3] [kg/m3] [%] c-7d 2320 1687 27.3 c-14d 2287 1687 26.2 c-28d 2234 1685 24.6 table 2. porosity evaluated from helium pycnometry. of m = 3.46 ± 0.02. lastly, the porosity ϕ0 was calculated with the known mean value of paramete m. all the calculated parameters are summarized in table 3. 5. conclusions in this paper, the porosity was measured by electrochemical impedance spectroscopy (eis) with the aid of helium pycnometry and evaluated using archie’s law. the resistivity of continuous conductive paths was evaluated from the impedance spectrum by the equivalent circuit (ec) and simplified equivalent circuit model (sec). the minimal differences between resistances rccp and r1 were found. hence, the capacitance of matrix c mat can be neglected in the sec model. thus, only the sec model is sufficient for evaluating the impedance spectra of pure portland cement paste. the value of archie’s index m was found to be almost identical between both ec and sec models and also for 7, 14, and 28 samples age. thus, the m parameter can be assumed as constant, which would allow porosity from eis measurement without the aid of another porosity measurement method. nevertheless, when time changes or supplementary material is added, the porous system also changes. therefore, further researches on cement-based materials are necessary. acknowledgements the works were supported by the czech science foundation (project 21-11965s) and the grant agency of the czech technical university in prague (sgs22/088/ohk1/2t/11). their support is gratefully acknowledged. 19 j. chalupová, j. němeček, v. hybášek et al. acta polytechnica ctu proceedings sample doh σ0 ϕ0,p rccp r1 σeff m ϕ0 ec sec ec sec ec sec m = 3.46 [%] [s/m] [%] [ω] [ω] [s/m] [s/m] [-] [-] [-] c-7d 65.5 13.08 27.3 201.3 204.1 0.147 0.145 3.459 3.470 27.3 c-14d 71.7 13.52 26.2 215.2 217.3 0.137 0.136 3.427 3.435 26.5 c-28d 76.8 13.90 24.6 274.5 276.8 0.108 0.107 3.467 3.473 24.5 table 3. the parameters necessary for the porosity calculation by archie’s law. references [1] y.-y. kim, k.-m. lee, j.-w. bang, s.-j. kwon. effect of w/c ratio on durability and porosity in cement mortar with constant cement amount. advances in materials science and engineering 2014:273460, 2014. https://doi.org/10.1155/2014/273460 [2] h. m. jennings. 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[20] a. e. a. hamami, j.-m. loche, a. aït-mokhtar. cement fraction effect on eis response of chloride migration tests. advances in cement research 23(5):233–240, 2011. https://doi.org/10.1680/adcr.2011.23.5.233 [21] b. j. christensen, t. coverdale, r. a. olson, et al. impedance spectroscopy of hydrating cement-based materials: measurement, interpretation, and application. journal of the american ceramic society 77(11):2789–2804, 1994. https: //doi.org/10.1111/j.1151-2916.1994.tb04507.x [22] g. e. archie. the electrical resistivity log as an aid in determining some reservoir characteristics. transactions of the aime 146(01):54–62, 1942. https://doi.org/10.2118/942054-g [23] r. nováková, m. kouřil, j. stoulil, et al. relation of corrosion simulators transport and modern concrete pore microstructure. in conference metal, pp. 1135–1140. 2015. [24] k. a. snyder, x. feng, b. d. keen, t. o. mason. estimating the electrical conductivity of cement paste pore solutions from oh-, k+ and na+ concentrations. cement and concrete research 33(6):793–798, 2003. https://doi.org/10.1016/s0008-8846(02)01068-2 [25] v. šmilauer, z. bittnar. microstructure-based micromechanical prediction of elastic properties in hydrating cement paste. cement and concrete research 36(9):1708–1718, 2006. https://doi.org/10.1016/j.cemconres.2006.05.014 21 https://doi.org/10.1680/adcr.2011.23.5.233 https://doi.org/10.1111/j.1151-2916.1994.tb04507.x https://doi.org/10.1111/j.1151-2916.1994.tb04507.x https://doi.org/10.2118/942054-g https://doi.org/10.1016/s0008-8846(02)01068-2 https://doi.org/10.1016/j.cemconres.2006.05.014 acta polytechnica ctu proceedings 40:15–21, 2023 1 introduction 2 electrochemical impedance spectroscopy 2.1 equivalent circuit models for concrete 2.2 porosity calculation from eis measurements 3 materials and methods 3.1 sample preparation 3.2 eis measurement 3.3 porosity measurement 4 results and discussion 5 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2023.40.0061 acta polytechnica ctu proceedings 40:61–68, 2023 © 2023 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague analysis of tunnel excavation based on linear dfn-fem modelling martin lebeda∗, petr kabele czech technical university in prague, faculty of civil engineering, department of mechanics, thákurova 7, 166 29 prague 6, czech republic ∗ corresponding author: martin.lebeda@fsv.cvut.cz abstract. simulations of tunnel excavations have to take into account the natural occurrence of joints and faults in the surrounding rock mass, which dominantly control its mechanical response. in this paper, we present work in progress toward 3d finite element analysis of excavation using equivalent rock-mass properties derived from stochastically generated discrete fracture networks (dfns). the equivalent stiffness is determined by volume averaging. presently, we solve the problem linearly for an incremental change of the stress state. the fracture’s stiffness is assumed to depend on the initial normal stress acting in direction normal to it. however, within the solved incremental step, we assume the fracture’s stiffness to be constant. this assumption is acceptable for small stress changes. since the fractures represented in the dfn model have preferred directions, the equivalent stiffness is anisotropic. keywords: fractured rock mass, discrete fracture network (dfn), finite element method (fem), averaging procedure. 1. introduction according to the current state of art in the field of rock mechanics, brittle structures (or fractures, including faults and joints) have the dominant effect on the overall mechanical response of the rock mass [1]. underground structures, such as tunnels, are often situated in rock masses with naturally existing fractures. since it is necessary to take into account the fractures’ influence on the deformation due to the changes in the stress field during excavation, estimating mechanical properties of the rock mass is fundamental for efficient and safe design of tunnel structures. the finite element method (fem) is a tool commonly used for simulating tunnel excavation. in the case of a highly fractured rock mass, it would not be effective and feasible to explicitly represent each distinct fracture in the finite element mesh. it is admissible to include only a few largest fractures or localized deformation zones using e.g. interface elements, while the rest of the fracture network is represented by so-called equivalent continuum. the overall effective properties can be determined by averaging or homogenization procedure [2]. in this paper, we use the averaging procedure proposed by oda et al. [3] to determine the effective stress-strain relation. as the overall properties of fractured rock mass depend on the size of the simulated area, it is suitable to include the concept of representative volume element (rve) or statistical volume element (sve) [4]. the discrete fracture network (dfn) method is one of possible ways to describe the fractures’ geometry while taking into account data obtained from structural-geological survey, e.g. [5–8]. one of the most used approaches are the stochastically generated dfns, which extrapolate the in-situ data by means of statistical probability distributions. in this paper, we present work in progress toward developing and implementing in computer code an approach that connects averaging procedure performed on 3d stochastic dfn with finite element analysis of tunnel excavation. at the present stage of development, we adopt some simplifying assumptions: (a) we assume that the rock mass in the whole domain of the finite element (fe) model has uniform equivalent effective mechanical properties. (b) the effective properties are derived by averaging [3] over a representative volume element (rve) of the rock mass, in which shapes, sizes, and orientations of individual fractures are modeled by means of stochastic dfn. the rve size is based on a previous study [9] and corresponds with the size of the fe model. (c) individual fractures are treated by means of the “parallel plate model” [3], which means that the deformation response of a fracture to normal and shear stress is represented by two parameters accounting for the fracture’s roughness, friction angle etc. as well as its size and normal stress acting on it. (d) the fe analysis is performed linearly for an incremental change of the stress state. the fracture’s stiffness is assumed to depend on the initial normal stress acting in direction normal to it. however, within the solved incremental step, we assume the stiffness to be constant. we are aware that some of the assumptions may be limiting the validity of the procedure. for example, 61 https://doi.org/10.14311/app.2023.40.0061 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en martin lebeda, petr kabele acta polytechnica ctu proceedings parameter value volumetric density p30 [1/m3] 2.0 power law exponent α [-] 3.4 power law minimum fracture size amin [m] 0.3 fisher distribution concentration factor κ [-] 106 fisher distribution mean vector [-] variable volume of the dfn model [m3] 100 × 100 × 100 rve volume [m3] 15 × 15 × 15 table 1. parameters of dfns. the size of rve (b) should be chosen not only based on the scatter of the overall properties [9], but also with respect to the gradient of the stress field in the fe model. or, the linear solution (d) is acceptable only for small stress changes. these limitations will be resolved in a follow-up work. 2. dfn model – description of fractures’ geometry quantitative information about fracturing of rock mass is often acquired by structural-geological mapping of fracture traces on “observation windows”, such as rock outcrops or tunnel walls. typically, the minimum recorded trace length is on the order of 10−1 m. due to the enormous number of corresponding fractures in the rock volume, on the one hand, and the limited information available on the observation windows, on the other hand, the geometry of the fracture network is described by stochastically generated dfn. in this approach, the size and orientation of the fractures are assumed to follow certain probabilistic distributions, whose parameters are identified so as to match (in a statistical sense) the traces’ sizes, directions and density within the observation windows. a dfn is then generated as a set of spatially distributed polygons, whose sizes and orientations respect the calibrated probabilistic distributions [8]. since each fracture is explicitly described in the dfn model, the network cannot be efficiently discretized by a 3d fe mesh. however, it provides a suitable basis for the use of an averaging procedure. the fracture network models used in this study were generated by the dfram software [8]. fractures’ centers are positioned in the model volume by using poisson random generation process. fractures’ sizes are controlled by power law distribution with parameters amin, which is the minimum fracture size (location parameter), and α, which is the law’s exponent (shape parameter). size of the fracture a is defined as the radius of the circle circumscribed to the fracture, while the fractures are modelled as squares. the orientation of fractures is determined by fisher distribution, with parameters µ, which is the mean unit normal vector of fractures, and κ, which is the concentration parameter. as the purpose of this study is to demonstrate the solution methodology, we adopt some simplifications while modelling the fracture networks. we use dfns consisting of one geological set of fractures. values of the power law distribution parameters are the same for all models and they are based on a previous study [9]. we consider that all fractures are nearly parallel, which is achieved by setting the concentration parameter of fisher distribution κ to a high value of 106. we will, however, consider fracture sets with different orientations, for which different mean normal vectors µ will be used (see subsection 6.1). the volumetric fractures density p30, which controls the total number of fractures in the model, was approximately set based on report [10]. parameters used in this study are listed in table 1. 3. overall stiffness of the fractured rock mass evaluated by averaging procedure the calculation of overall effective parameters of fractured rock mass is based on the volume averaging relation proposed by oda et al. [3]: εij = 1 e [ (1 + ν)δikδjl − νδij δkl + ( 1 kn − 1 ks ) fijkl + 1 4ks (δikfjl + δjkfil + δilfjk + δjlfik) ] σkl =cijklσkl, (1) where e and ν are young’s modulus and poisson’s ratio of the intact rock, respectively, kn and ks are nondimensional parameters related to the fracture’s normal and tangent stiffness, respectively, δij is kronecker’s delta, and fij and fijkl are so-called second and fourth rank crack tensors, respectively: fij = 1 v m∑ p=1 s(p)l(p)n (p) i n (p) j , (2) fijkl = 1 v m∑ p=1 s(p)l(p)n (p) i n (p) j n (p) k n (p) l . (3) here v is the volume of the domain over which averaging is performed, s(p) is the area of p-th fracture 62 vol. 40/2023 analysis of tunnel excavation based on linear dfn-fem modelling figure 1. normal-stress vs. fracture-closure diagram [12] and schematic indication of the tangent stiffness evaluation. inside the sampling volume, l(p) is a typical size of the fracture equal to a diameter of a circle with the same area and n(p)i are the components of a unit vector normal to the fracture. note that the typical size l is assumed as a property of the fracture independent on the size of the sampling volume and it is calculated using the original area of the fracture generated in stochastic dfn. the calculation procedure of parameters kn and ks is described in detail in our previous work [11]. in brief, the parameters are expressed as: kn = 3π 8 + κn · l e , (4) ks = 3π 8 + κs · l e , (5) where κn and κs are the normal and shear stiffness (stress/relative displacement) of a fracture and l is the fracture’s size. the normal stiffness parameter κn is to be obtained experimentally. it depends on whether the fracture undergoes loading or unloading and it also varies with pressure acting on it. for the sake of simplicity we determined κn as tangent stiffness by piece-wise linear approximation of the unloading path of the experimental normal-stress vs. fracture-closure diagram (figure 1), which was presented in [12]. the intervals of normal stress σn and corresponding tangent stiffness parameters κn and kn are listed in table 2. compressive stress is in this study considered with a negative sign. the shear stiffness of fractures is determined using the formula presented in [13] with parameters listed in table 3: κs = 100 l · |σn| · tan ( j rc · log10 ( j cs |σn| ) + ϕr ) , (6) where j rc is joint roughness coefficient, j cs is joint wall compressive strength and ϕr is residual friction angle. normal stress |σn| κn kn [mpa] [ mpam ] [-] 0.0–2.5 75 472 1.308 2.5–5.0 210 526 1.540 5.0–7.5 277 778 1.656 7.5–10.0 625 000 2.253 10.0–12.5 781 250 2.522 12.5–15.0 892 857 2.714 15.0–20.0 1 250 000 3.328 20.0–25.0 1 250 000 3.328 25.0–30.0 1 315 789 3.442 30.0–33.0 2 790 698 5.979 table 2. fracture’s tangent stiffness (in direction normal to the fracture). parameter value joint roughness coefficient j rc [-] 6.55 joint wall compressive strength j cs [mpa] 90.0 residual friction angle ϕr [deg] 20.0 table 3. fracture’s shear stiffness parameters. the described constitutive model has been implemented by means of python api as user-defined material in the open-source fe code oofem [14]. it is noted that equation (3) represents a compliance relation, while in fem, the stiffness tensor is necessary. thus, the compliance tensor is first calculated using equation (3) and then it is numerically inverted to obtain stiffness. 4. representative volume element as we discussed in the previous section, the overall mechanical properties of the jointed rock mass are evaluated by volume averaging of the fracture network. the size of the dfn sub-domain, over which the averaging is performed (v in equations (2) and (3)), should be chosen large enough to guarantee that the evaluated overall properties are the same for any stochastic realization of the dfn. then, the subdomain can be called a representative volume element. based on a previous study [9], we use rves with dimensions of 15 × 15 × 15 m. for this size the coefficient of variation (cov) of apparent moduli evaluated for 10 realizations of dfn (with similar parameters as here) was less than 20 %. furthermore, it was shown in [9], that further increasing the sampling volume size did not lead to any significant reduction of the moduli cov. to eliminate potential border effects, the rves are obtained by cropping much larger dfn, which is generated in the volume of 100 × 100×100 m. 63 martin lebeda, petr kabele acta polytechnica ctu proceedings (a). state prior to excavation. (b). state after tunnel excavation. figure 2. finite element model. 5. 3d fe modelling of tunnel excavation to demonstrate the use of the dfn-based constitutive model, we numerically simulate by fem the excavation of a fictitious tunnel in rock at the depth of 550 m. the tunnel has a circular cross-section with the diameter of 3.0 m. the tunnel is modelled in a cubic body, representing the surrounding rock, with side size of 15 m (figure 2). thus, the distance between the tunnel and the boundary of the model is at least twice the diameter of the construction. the coordinate system is introduced with the vertical axis z oriented upwards. the edges of model are parallel with the coordinate axes and the tunnel is parallel with the y-axis. the finite element mesh consists of 3d hexahedrons and it was created with salome software [15]. the simulation is performed in two steps. the first step (time t = 0) represents the initial state prior to excavation – figure 2a. the displacement of the rock body is constrained by statically determinate supports and surface tractions, corresponding to the initial geostatic pressure, are applied on all sides of the body. the initial stress state is idealized as uniform within the modeled domain. the initial stress tensor is based on field measurements from rožná mine [16]. using data from vertical boreholes it was found that the directions of the principal stresses are vertical and horizontal. the vertical principal stress is calculated by equation (7): sv = −ρ · g · h, (7) where ρ is volumetric weight, g is gravity acceleration and h is depth beneath the ground level. the parameters for calculation of the vertical principal stress (table 4) are adopted from [17]. as the major and minor horizontal principal stresses, sh and sh, respectively, we use the average of the values measured on different levels of the mine. the values of the parameter value major horizontal principal stress sh [mpa] −22.8 minor horizontal principal stress sh [mpa] −14.6 volumetric weight ρ [ kg 3 m ] 28 000 depth of the mine beneath the ground level [m] 550 vertical principal stress sv [mpa] −15.4 table 4. the initial stress state. initial stress are listed in table 4. orientation of the model is set so, that the tunnel axis is parallel with the minor horizontal principal stress, which means that the direction of sh corresponds to the global axis y. it should be noted that, even though the stress field is uniform throughout the analyzed domain in the first step, the normal traction σn acting on each fracture in the dfn is generally different due its different direction. this traction, in turn, affects the fractures’ normal and shear stiffness, as seen in equation (6) and table 2, which must be taken into account in the averaging process. in the second calculational step (time t = 1), the excavation of the tunnel is simulated by removing the finite elements inside the tunnel space while keeping the boundary conditions on the outer surface of the rock body unchanged. this creates a new tractionfree boundary on the tunnel walls, which results in deformation and change of stress state in the rock. we should remark that, at the present stage of progress, the problem is solved as incrementally linear. that is, the second step is solved with constant tangent stiffnesses of the fractures, which correspond to the initial stress state. 64 vol. 40/2023 analysis of tunnel excavation based on linear dfn-fem modelling figure 3. horizontal convergence ∆ux calculated for different dfn models. figure 4. vertical convergence ∆uz calculated for different dfn models. 6. case examples 6.1. fracture network models to demonstrate the proposed simulation procedure of 3d underground structure excavation, we use several simple examples. in all of the examples, the fracture network consists of a set of nearly parallel fractures with the same dfn parameters, except the mean normal vector. in particular, the simulations are performed on four types of dfn models, in which fractures are: perpendicular to axis x/y/z and at an angle of 45 degrees between negative x axis and positive z axis, see table 1. even though the size of the rve for averaging of the rock properties was rationally selected, 3 different stochastic realizations of dfn models with different random seed value were generated for verification. thus, in total we work with 12 dfn models. 6.2. results and discussion 6.2.1. displacements figure 3 and figure 4 present the calculated values of the horizontal and vertical convergence for each dfn case. the tunnel convergence is evaluated as the relative displacement between points on the tunnel walls caused by the tunnel excavation (that is, the incremental displacement between the final displacement at time t = 1 and the initial displacement at t = 0). positive values of the convergence correspond to the closing of the tunnel walls. the points at which figure 5. location of reference points and lines used to report results. the displacement was recorded are pictured in figure 5. figure 6 through figure 9 show the incremental displacement of the model boundaries including the excavated tunnel. considering that the initial vertical pressure sv is about 2⁄3 of the major horizontal pressure sh and 65 martin lebeda, petr kabele acta polytechnica ctu proceedings that sh acts perpendicularly to the tunnel axis, the displacements of the tunnel walls qualitatively correspond with expectations: • for a fracture set perpendicular to tunnel axis, i.e. perpendicular to the axis y, the effect of the fractures on the convergence is rather small and the convergence is the least among all the analyzed cases (the second group of bars in figure 3 and figure 4). consistently with the initial stress state, the vertical convergence is smaller than the horizontal one, which is also obvious from the deformed shape of the tunnel cross-section (figure 6). • referring to the cases with vertical and horizontal fractures parallel with the tunnel axis (i.e. perpendicular to axis x and perpendicular to axis z, respectively), it is seen in figure 3 and figure 4 that the convergence in the direction normal to the fractures significantly increases in comparison with the previous case. the convergence as well as the shape of the cross-sectional distortion reflect the fractures-induced anisotropy. in the rock mass with vertical fractures, the higher initial horizontal stress results in horizontal convergence, which is significantly larger than the vertical one (figure 7). on the other hand, when the fractures are horizontal, the resulting horizontal and vertical convergences are close to each other (figure 8). • finally, when the fractures are inclined at 45 degrees, the horizontal and vertical convergences fall between those observed with the vertical and horizontal sets. the axes of the distorted shape of the tunnel (figure 9) are close to, but do not align with the fractures, which can be attributed to the difference of the initial horizonal and vertical stress and normal and tangential displacement occurring on the fractures. figure 3 and figure 4 also indicate that the larger components of convergence, which are strongly affected by the fractures, exhibit notable variation among the different stochastic realizations of the dfn. this finding suggests that the employed sampling volume was still not large enough to meet the criteria as an rve. 6.2.2. stress figure 10 and figure 11 show the distribution of normal stresses along the horizontal and vertical lines indicated in figure 5. the results are reported for one realization of the dfn model with fractures at the angle of 45 degrees. both plots show the uniform initial stress state at time t = 0 and the stress distribution after excavation at t = 1. the difference between the lines corresponds to the incremental change of stress due to excavation. the plots show that, as the tunnel is excavated and a new traction-free boundary is introduced at the tunnel wall, the radial pressure correctly decreases from the initial uniform state and tends to zero at the tunnel wall. on the other hand, (a). vertical section through the dfn. (b). deformed shape of the tunnel (red). figure 6. model with fractures perpendicular to y axis. (a). vertical section through the dfn. (b). deformed shape of the tunnel (red). figure 7. model with fractures perpendicular to x axis. (a). vertical section through the dfn. (b). deformed shape of the tunnel (red). figure 8. model with fractures perpendicular to z axis. (a). vertical section through the dfn. (b). deformed shape of the tunnel (red). figure 9. model with fractures at the angle of 45 degrees between negative x axis and positive z axis. 66 vol. 40/2023 analysis of tunnel excavation based on linear dfn-fem modelling figure 10. model with fractures at the angle of 45 degrees: distribution of normal stresses σx and σz along horizontal line a-a’ (figure 5). the pressure in the circumferential direction increases, which is also to be expected. note that fractures are distributed at an angle around the tunnel, so some experience an increase of pressure (loading) while pressure on others decreases (they unload). the stress increment on fractures in the close vicinity of the tunnel may be also quite large. therefore, for some fractures, using the constant tangent stiffness estimated from the unloading path of the stress-closure diagram (see section 3 and figure 1) might have not been accurate. 7. conclusions the 3d fem simulation of tunnel excavation with rock mass parameters evaluated by averaging procedure based on dfn has been examined. this paper has been focused on the demonstration of the solution methodology. the results allow us to draw the following conclusions: • deformed shapes calculated by fe simulations are consistent with the orientation of fractures in dfns and the initial stress state. largest tunnel convergence is observed in the directions close to the least overall stiffnesses of the fractured rock mass. • some difference in the displacements calculated with different stochastic realizations of the dfns was observed. the appropriate size of the sampling volume should be further investigated and if rve cannot be found, the concept of sve might be adopted. • the largest changes in the stress state after the tunnel excavation were observed close to tunnel, figure 11. model with fractures at the angle of 45 degrees: distribution of normal stresses σx and σz along vertical line b-b’ (figure 5). which is in agreement with the expectations. on the other hand, on the outer boundaries of the models, the stresses after excavation were slightly different from the initial state, which indicates that a larger domain of the model is required to eliminate the boundary effect. • considering the observed stress variations after the tunnel excavation, an incremental nonlinear calculation with fractures’ stiffness updated according to the actual stress state and loading-unloading may be necessary for a more accurate prediction of the tunnel convergence. • although it was not elaborated in the present work, the demonstrated analysis method provides information about the changes of the pressure acting on the fractures, which affect their hydrogeological properties. the proposed methodology, therefore, may provide valuable inputs for hydrogeological simulations of ground water flow and contaminants’ transport. acknowledgements this paper was financially supported by czech technical university in prague under sgs project no. sgs22/030/ohk1/1t/11. references [1] l. jing. a review of techniques, advances and outstanding issues in numerical modelling for rock mechanics and rock engineering. international journal of rock mechanics and mining sciences 40(3):283–353, 2003. https://doi.org/10.1016/s1365-1609(03)00013-3 67 https://doi.org/10.1016/s1365-1609(03)00013-3 martin lebeda, petr kabele acta polytechnica ctu proceedings [2] m. hori, s. nemat-nasser. on two micromechanics theories for determining micro–macro relations in heterogeneous solids. mechanics of materials 31(10):667–682, 1999. https://doi.org/10.1016/s0167-6636(99)00020-4 [3] m. oda, t. yanabe, y. ishizuka, et al. elastic stress and strain in jointed rock masses by means of crack tensor analysis. rock mechanics and rock engineering 26(2):89– 112, 1993. https://doi.org/10.1007/bf01023618 [4] x. yin, w. chen, a. to, et al. statistical volume element method for predicting microstructure–constitutive property relations. computer methods in applied mechanics and engineering 197(43-44):3516–3529, 2008. https://doi.org/10.1016/j.cma.2008.01.008 [5] w. s. dershowitz, p. wallmann, s. kindred. discrete fracture modelling for the stripa site characterization and validation drift inflow predictions, stripa-tr-91-16, 1991. skb stockholm. https://www.researchgate.net/profile/william_ dershowitz/publication/266334817_discrete_ fracture_modelling_for_the_stripa_site_ characterization_and_validation_drift_inflow_ predictions/links/563bb3ea08ae34e98c47cfaf.pdf [6] r. munier. statistical analysis of fracture data, adapted for modelling discrete fracture networks-version 2, skb r-04-66, 2004. skb svenskkärnbränslehantering ab swedish nuclear fuel and waste management co. http: //skb.se/upload/publications/pdf/r-04-66.pdf [7] c. darcel, p. davy, o. bour, j.-r. d. dreuzyr. discrete fracture network for the forsmark site, skb rapport r-06-79, 2006. skb svenskkärnbränslehantering ab swedish nuclear fuel and waste management co. http://www.iaea.org/inis/collection/ nclcollectionstore/_public/38/013/38013872.pdf [8] p. kabele, et al. mathematical modeling of brittle fractures in rock mass by means of the dfn method – final report súrao tz 286/2018/eng, 2017. súrao radioactive waste repository authority. [9] m. lebeda, p. kabele. the effect of sampling volume size on the apparent stiffness of jointed rock mass. app 34:38–42, 2022. https://doi.org/10.14311/app.2022.34.0038 [10] l. gvoždík, et al. transport of radionuclides from deep geological repository/testing of conceptual and numeric models – final report súrao tz 463/2020/eng, 2020. súrao radioactive waste repository authority. [11] m. lebeda. určení tuhosti porušeného horninového masivu s využítím dfn modelů. master’s thesis, czech technical university in prague, faculty of civil engineering, department of mechanics, 2020. [12] s. c. bandis, a. c. lumsden, n. r. barton. fundamentals of rock joint deformation. international journal of rock mechanics and mining sciences & geomechanics abstracts 20(6):249–268, 1983. https://doi.org/10.1016/0148-9062(83)90595-8 [13] n. barton, v. choubey. the shear strength of rock joints in theory and practice. rock mechanics 10(1-2):1– 54, 1977. https://doi.org/10.1007/bf01261801 [14] b. patzák. oofem project home page, 2000. [2022-10-13]. http://www.oofem.org [15] edf, cea. salome, 2001. [2022-10-17]. https://www.salome-platform.org/ [16] z. bukovská, a kol. získání dat z hlubokých horizontů dolu rožná, tz 464/2020, 2020. súrao správa úložišť radioaktivních odpadů. [17] k. souček, m. vavro, l. staš, et al. komplexní geologická charakterizace prostorů pvp bukov – část ii geotechnická charakterizace, závěrečná zpráva súrao zz 221/2018, 2017. súrao správa úložišť radioaktivních odpadů. 68 https://doi.org/10.1016/s0167-6636(99)00020-4 https://doi.org/10.1007/bf01023618 https://doi.org/10.1016/j.cma.2008.01.008 https://www.researchgate.net/profile/william_dershowitz/publication/266334817_discrete_fracture_modelling_for_the_stripa_site_characterization_and_validation_drift_inflow_predictions/links/563bb3ea08ae34e98c47cfaf.pdf https://www.researchgate.net/profile/william_dershowitz/publication/266334817_discrete_fracture_modelling_for_the_stripa_site_characterization_and_validation_drift_inflow_predictions/links/563bb3ea08ae34e98c47cfaf.pdf https://www.researchgate.net/profile/william_dershowitz/publication/266334817_discrete_fracture_modelling_for_the_stripa_site_characterization_and_validation_drift_inflow_predictions/links/563bb3ea08ae34e98c47cfaf.pdf https://www.researchgate.net/profile/william_dershowitz/publication/266334817_discrete_fracture_modelling_for_the_stripa_site_characterization_and_validation_drift_inflow_predictions/links/563bb3ea08ae34e98c47cfaf.pdf https://www.researchgate.net/profile/william_dershowitz/publication/266334817_discrete_fracture_modelling_for_the_stripa_site_characterization_and_validation_drift_inflow_predictions/links/563bb3ea08ae34e98c47cfaf.pdf http://skb.se/upload/publications/pdf/r-04-66.pdf http://skb.se/upload/publications/pdf/r-04-66.pdf http://www.iaea.org/inis/collection/nclcollectionstore/_public/38/013/38013872.pdf http://www.iaea.org/inis/collection/nclcollectionstore/_public/38/013/38013872.pdf https://doi.org/10.14311/app.2022.34.0038 https://doi.org/10.1016/0148-9062(83)90595-8 https://doi.org/10.1007/bf01261801 http://www.oofem.org https://www.salome-platform.org/ acta polytechnica ctu proceedings 40:61–68, 2023 1 introduction 2 dfn model – description of fractures' geometry 3 overall stiffness of the fractured rock mass evaluated by averaging procedure 4 representative volume element 5 3d fe modelling of tunnel excavation 6 case examples 6.1 fracture network models 6.2 results and discussion 6.2.1 displacements 6.2.2 stress 7 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0470 acta polytechnica ctu proceedings 38:470–477, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague use of energy profile indicators to determine the expected range of heating energy consumption tobias loga∗, britta stein, guillaume behem iwu – institut wohnen und umwelt, institute for housing and environment, rheinstraße 65, de-64295 darmstadt, germany ∗ corresponding author: t.loga@iwu.de abstract. a small set of query variables designed to collect information about the energy-related features of residential buildings is presented. these “energy profile indicators” include information about those visible characteristics of a building which have a notable impact on its energy performance and are simple to assess. the queries are an interesting source for a rough energy performance calculation for single buildings as well as for housing portfolios or housing stocks. a method has been developed to transform the energy profile indicators into input data for a physical calculation model. it consists of procedures to estimate the envelope area, u-values, and efficiency values of the heat supply system. to all model input variables an uncertainty is assigned. if information from a query is not available, the model input is set to a state representing the average building stock and the uncertainty of this quantity is adjusted to a value reflecting the variance in the stock. the resulting uncertainty of the calculated energy use is determined. examples of the application of the method are given to show the influence of different unknown quantities including occupant behaviour. experiences on the coherence with metered consumption are reported. keywords: monitoring indicators, building stock, surveys, statistical evaluation, realistic physical model, uncertainties, target/actual comparison. 1. introduction various activities in the building sector aim at gathering building data, executing energy performance calculations and assessing the metered energy consumption, often in the context of energy performance certificates (epcs). however, huge challenges are faced when cross-section and longitudinal analyses are needed to get an overall view of the refurbishment progress and energy consumption of housing stocks. this is not only due to data gaps but also to the used calculation methods and software applications focussing on official proof of requirements and rating. the detailedness and complexity of input data, recurrent revisions of the regulatory framework, but also the lack of software applications to store, calculate and analyse multi-building data are barriers for aggregation and statistical evaluation. the idea of the concept presented in this paper is to tackle these challenges by introducing an additional level of energy assessment addressing evaluation purposes. the overall objectives are: • facilitate the collection and maintenance of information about energy-related features of housing stocks, allowing for an assessment of the overall state of refurbishment, annual refurbishment rates and for a comparison with milestones on the refurbishment path. • provide reliable estimates of the energy consumption of single buildings and of housing stocks for different levels of insulation and various heat supply system types to be used for prognoses in the context of refurbishment scenarios. • enable a target/actual comparison of the consumption after implementation of refurbishment measures for single buildings and also for housing stocks. in order to reach these aims, a methodical framework has been developed within the research project mobasy, consisting of two parts [1]: (1.) a set of monitoring indicators representing those visible characteristics of buildings that have the most important effect on its energy performance, suitable for use in questionnaires and quick on-site inspections. (2.) an energy performance calculation using the mentioned monitoring indicators as input data and providing a realistic bandwidth of the expected energy consumption, reflecting the uncertainty of the data acquisition (uncertain or unavailable information) and of the model input data (thermal properties of materials and components, occupant behaviour, climate) depending on the specific field of application (assessment of past consumption or prognosis). this paper gives an overview of the monitoring indicators (“energy profile indicators”), and of the principles of the uncertainty assessment attached to the realistic energy performance calculation (“simple combined physics-probability model”). the influence of 470 https://doi.org/10.14311/app.2022.38.0470 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 use of energy profile indicators to determine the expected range . . . characteristics of indicators (query variables) building size and form living space, number of attached buildings, number of full storeys, heating situation in attic and basement, number of dwellings thermal properties of the opaque envelope (categories: roof, ceiling, wall, floor) building: year of construction; per envelope category: existent insulation upgrade, year of insulation upgrade, insulation thickness + covered fraction thermal properties of windows (two types) per window type: number of panes, low-e coating existent, type of window frame, year of installation, fraction of the second window type, if applicable heat supply system (heating / dhw) heat generators (type, specification, year of installation), thermal solar system, heat storages (type, location inside or outside of thermal envelope), heat distribution (type, insulation level, location inside or outside of thermal envelope) further systems ventilation system (with / without heat recovery), pv system (with / without electrical storage) table 1. overview of energy profile indicators for residential buildings. the origin and detailedness of data on the uncertainty of the estimated energy consumption is illustrated for a singleand a multi-family house. furthermore, results are presented from the application of the method to a housing stock sample, providing statistical information on the energy-related state, on the estimated and actual consumption. 2. energy profile indicators a set of “energy profile indicators” has been developed in [2], then continuously further refined and applied in several projects for portfolio assessment and quality assurance (examples: [3, 4]). the indicators have also been used in house owner surveys in 2009 [5] and 2016 [6] on the basis of random sampling1 to get an image of the refurbishment state and progress of the german residential building stock. table 1 gives an overview of the query variables. these indicators represent the physical characteristics of a building that have the biggest impact on its energy use for heating and domestic hot water (dhw). the variables are designed to gather information by asking (technically informed) building owners and by executing on-site inspections of residential buildings. for these purposes, a two-page questionnaire [7] and a structured data table [8] have been developed. energy profile indicators provide information that can be used to estimate the energy consumption of a building. for this purpose, a transformation to a physical model including an uncertainty assessment has been developed. 3. simple combined physics-probability model there are different approaches to quantify the uncertainties of the building energy assessment [9]. the 1the 2016 evaluation is based on about 17 000 evaluable questionnaires (18 % of ca. 92 000 queries sent to a random sample of german house owners) [6]. concept presented here uses forward uncertainty, that tries to quantify the uncertainty in the calculation outputs propagated from uncertain input variables through mathematical models. due to the stationary, quasi-linear energy calculation model, a very simple, non-sampling probabilistic method can be employed, assuming normal distributions and applying the gaussian error propagation law. 3.1. principle the principle of the indicator-based energy performance calculation is shown in figure 1. the first step is a transformation of energy profile indicators to input variables of the physical model. as a result, a “calculation value” is provided for each quantity, directly used by the energy performance calculation (“model input variables”), and an “uncertainty range” is attributed, which represents the span of values that might include the actual (unknown) real value. preferably, the parameters of the transformation should be derived from empirical data of large building stock samples. however, for certain quantities extended surveys or large measurement campaigns are not available. in these cases, the calculation value and uncertainty range can only be estimated. a pragmatic way to do this is to ask an expert: which is the highest, which is the lowest value that might be present and still be seen as “not unusual”? the average of both is the calculation value, half the difference between both represents an indicator of the uncertainty. as an important means for transparency and inducement for continuous improvement of the method, the derivation of calculation values and uncertainties must be documented for each variable. the calculation values are used as input for the physical model to determine the theoretical energy consumption. in a separate calculation, the uncertainty of the calculated energy consumption is determined by combining the effect of the uncertainties of all input 471 t. loga, b. stein, g. behem acta polytechnica ctu proceedings figure 1. principle of the combined physics-probability model based on energy profile indicators. values. gaussian error propagation law (square root of the added-up squares of the uncertainties caused by each single variable) is used for this purpose. since most of the probability density distributions of input variables are assumed to be symmetrical and the energy balance procedure used (seasonal method) is mainly linear, the systematic deviations produced by this simplification are assumed to be small. the result of this combined probability-physics model is an expectation value and an expectation range of the energy use for heating and dhw. 3.2. systematic uncertainty assessment the uncertainty of the energy performance calculation is in principle depending on the type of data source and on the completeness of information (see figure 2). type of data source: information provided by building owners is a useful data source, however inspections can provide additional information and thus reduce the uncertainty. more trustworthy are design data from refurbishment planning (including issued epcs), at best if the implementation is secured by quality assurance (qa). completeness of information: information can be missing from all mentioned data sources (even in a quality assured planning process). when typical or average values from the building stock are used instead the uncertainty of the calculation increases. this global principle has been translated into an algorithm: for each model input variable five uncertainty classes a to e are defined with rules for selection and uncertainty values to be used. as an example, table 2 shows the uncertainty assessment of the insulation thickness, applied separately for all opaque envelope categories. analogue definitions are available for the other input variables2. 2these are: envelope area, u-values of original constructions, insulated fractions by envelope type, thermal conductivities figure 2. influences on the uncertainty of an energy performance calculation. the method presented in this paper is focussing on data from building owners and from on-site inspections, associated with uncertainty classes c and d. these data sources facilitate the energy assessment of a large number of buildings – as a start from files and records – updates can be provided in the course of insulation by envelope type, u-value of windows, thermal bridging surcharge, heating degree days, annual solar radiation, effective passive solar aperture, internal heat load, efficiency of ventilation heat recovery, dhw heat need, efficiency of heating system, efficiency of dhw system. tabled values and explanations in english language are displayed at the end of the supplemental document [10]. a detailed description of the algorithms for a selection of variables is given in annex d of [1] (german language). 472 vol. 38/2022 use of energy profile indicators to determine the expected range . . . uncertainty description uncertainty of class insulation thickness∗ relative absolute a insulation thickness from design data, quality assured 5 % 0.5 cm b insulation thickness from design data 10 % 1 cm c insulation thickness determined by inspection on-site / record or statement of building owner 25 % 2 cm d insulation available; insulation thickness unknown; use of average values from the building stock differentiated by implementation period 40 % 5 cm e insulation thickness and year of implementation unknown; use of average values from the building stock (all buildings) 50 % 8 cm ∗ relevant for the calculation is the minimum in cm, derived from the relative and absolute uncertainty. table 2. example of uncertainty classification for an input variable (insulation thickness). of on-site investigations. even if input is missing for some indicators (class e), the estimation of the actual energy use of the sample or portfolio can be commenced – with corresponding larger expectation ranges. if available, design data with or without quality assurance (classes a or b) can also be integrated. the improvement achieved by gradually adding reliable data from inspections and from quality assured refurbishments will result in narrowed ranges of the expected energy use. 3.3. mechanisms of the energy performance calculation based on energy profile indicators following the above-described principles, a coherent method for an energy performance calculation by use of energy profile indicators has been developed and implemented in form of an excel-workbook and an r script (description of the method in [1]). core aspects are the estimation of the thermal envelope area [11] and of the u-values [12]. the energy performance calculation is based on the tabula seasonal method [13], using climate data differentiated by postcode: in the first instance, the physical model is calculated by use of the local long-term average climate (of the past 20 years). for performing target/actual comparisons for specific years, calculated heat losses and gains are calibrated by the actual temperatures and solar radiation of the metering period. 4. showcase: influence of different parameters on the uncertainty of the estimated consumption for a single building, the described method allows a prognosis of the energy consumption, providing an estimation value (result of the physical model) and an estimation range (result of the uncertainty assessment). the actual consumption is expected to be found within the estimation range, however in rarer cases it can also go beyond that. table 3 shows examples of expectation ranges for a single-family and a multi-family house (sfh, mfh), depending on the state of refurbishment, on the data acquisition type and on the utilisation of the buildings. these are results from the combined physicsprobability model with uncertainty assessment as described above. the examples are intended to illustrate the uncertainty of the consumption estimate based on energy profile indicators (house owner statement) as opposed to those based on design data, considering particularly the influence of unknown occupant behaviour. for the variants 1.1 to 1.3 the calculation is assumed to be based on energy profile indicators, with input from house owners. for the single-family house, the relative uncertainty of the calculated energy use for heating and dhw is ±35 % before (var. 1.1) and ±72 % after a deep refurbishment with passive house components (var. 1.2). the relative increase is mainly caused by the influence of occupant behaviour. however, the state of the building is much better defined after refurbishment, so the absolute uncertainty caused by the envelope is only ±25 kwh/(m2a) compared to ±52 kwh/(m2a) before refurbishment. if the real climate conditions of the climate zone are used, the uncertainty of the energy demand caused by the climate is reduced from ±26 % (var. 1.2) to ±19 %, (var. 1.3), having only a small effect on the total uncertainty. in variant 1.4, design data from the planning of the refurbishment (detailed calculation of envelope area, of u-values, and of supply system performance) and long-term climate data are used. this is the typical pattern of energy performance certificates, issued in the context of modernisation. it reduces the total uncertainty from about ±70 to ±60 %. the effect is rather small due to the predominating large uncertainties of utilisation and climate. variant 1.5 shows a situation where the design data are quality-assured (update of information about actual implementation 473 t. loga, b. stein, g. behem acta polytechnica ctu proceedings e xe m pl ar y bu ild in g sf h (s in gl efa m ily ho us e) [1 95 8– 19 68 ] m f h (m ul ti -f am ily ho us e) [1 95 8– 19 68 ] st at e of re fu rb is hm en t u nr ef ur bi sh ed r ef ur bi sh ed (p as si ve ho us e co m po ne nt s) u nr ef ur bi sh ed r ef ur bi sh ed (p as si ve ho us e co m po ne nt s) t yp e of da ta so ur ce e x po st (h ou se ow ne r st at em en t) d es ig n da ta d es ig n da ta + q a e x po st (h ou se ow ne r st at em en t) d es ig n da ta d es ig n da ta + q a c lim at e lo ng -t er m av er ag e sp ec ifi c ye ar lo ng te rm av er ag e sp ec ifi c ye ar lo ng -t er m av er ag e sp ec ifi c ye ar lo ng te rm av er ag e sp ec ifi c ye ar o cc up an t b eh av io ur un kn ow n de ta ile d m ea su re m en ts ∗ un kn ow n de ta ile d m ea su re m en ts ∗ v ar ia nt 1. 1 1. 2 1. 3 1. 4 1. 5 2. 1 2. 2 2. 3 2. 4 2. 5 calculatedenergy use(naturalgas) [kwh/(m2a)] e ne rg y us e∗ ∗ • h ea ti ng 20 3 50 50 50 50 15 6 31 31 31 31 • d h w 25 9 9 9 9 26 15 15 15 15 su m 22 8 59 59 59 59 18 3 45 45 45 45 u nc er ta in ty ∗ ∗ ∗ • e nv el op e ± 23 % ± 42 % ± 42 % ± 15 % ± 7 % ± 17 % ± 29 % ± 29 % ± 12 % ± 5 % • sy st em ± 9 % ± 10 % ± 10 % ± 7 % ± 5 % ± 9 % ± 12 % ± 12 % ± 10 % ± 6 % • u ti lis at io n ± 19 % ± 52 % ± 52 % ± 52 % ± 13 % ± 4 % ± 14 % ± 14 % ± 14 % ± 4 % • c lim at e∗ ∗ ∗ ∗ ± 16 % ± 26 % ± 19 % ± 26 % ± 13 % ± 16 % ± 33 % ± 24 % ± 30 % ± 10 % t ot al ± 35 % ± 72 % ± 70 % ± 60 % ± 20 % ± 26 % ± 47 % ± 42 % ± 37 % ± 14 % ∗ in do or te m p er at ur es an d w in do w s op en in g ha ve b ee n m ea su re d, e. g. in a re se ar ch pr oj ec t (a ss um pt io n fo r th e sh ow n ca se : bu ild in g ha s ut ili sa ti on cl os e to th e av er ag e of al l bu ild in gs ); ho w ev er , in te rn al he at so ur ce s an d ex te rn al sh ad in g ha ve no t b ee n m ea su re d or m od el le d in de ta il. ∗ ∗ n at ur al ga s, in kw h/ (m 2 a) ; ca lc ul at io n w it h si m pl e pr ob ab ili ty -p hy si cs m od el . ∗ ∗ ∗ r el at iv e un ce rt ai nt y, re la te d to en er gy us e fo r he at in g an d d h w . ∗ ∗ ∗ ∗ in cl ud in g un ce rt ai nt y of ex te rn al sh ad in g (a ss um pt io n fo r al l va ri an ts : n o de ta ile d as se ss m en t of th e re du ct io n of pa ss iv e so la r ga in s ca us ed by su rr ou nd in g tr ee s an d bu ild in gs or by us e of sh ut te rs et c. ). table 3. estimation of the actual energy consumption before and after deep refurbishment for two example buildings and assigned uncertainty, depending on the type of data source and measurements (a pdf demonstrating the calculation flow for the 10 building variants can be viewed in the supplemental document mentioned above [10]). 474 vol. 38/2022 use of energy profile indicators to determine the expected range . . . during refurbishment) and where the utilisation and the climate are measured in detail, for example in the framework of a research project. in this case, the uncertainty of the consumption estimate can be reduced to ±20 %3. for the multi-family house with 24 flats (var. 2.1 to 2.5), the effect of the utilisation uncertainty is significantly smaller because of stochastic effects (compensation between flats with high and low intensity of usage). it may be noted that a very deep refurbishment has been assumed here, resulting in very low energy consumption and causing large relative uncertainties. however, the absolute uncertainty of about ±40 kwh/(m2a) for sfh (var. 1.3) and of ±20 kwh/(m2a) for mfh (var. 2.3) is not changing much for less ambitious refurbishments, resulting in smaller relative uncertainties (see results for the building sample presented in the following chapter). if design data are not available for such a deep refurbishment, the thermal envelope’s uncertainty is predominated by the unknown effects of thermal bridging. if a characterisation is possible (for example as a result of on-site inspection: no relevant constructive weaknesses), the uncertainty of the energy use caused by the envelope can be reduced from ±29 % to ±16 % (mfh, var. 2.3), the total uncertainty reduced from ±42 % to ±34 %. these uncertainties of consumption estimates are valid with view at a single building. however, if large building samples or stocks are considered, the uncertainty of the predicted total consumption (and thus the uncertainty of the average) is reduced to 1/10 of the shown values for 100 buildings and to 1/100 for 10 000 buildings. in summary, the expectation ranges of the energy consumption determined by quantifying input uncertainties seem rather large, especially for single-family houses. this is not only true for the calculation based on energy profile indicators but also for the epc calculation based on design data. for both, the predominating factors determining the uncertainty of the prognosis of an annual consumption are unknown occupant behaviour and climate conditions. 5. application on a sample of apartment blocks from housing companies in the following, the application of the combined probability-physics model on a housing sample from three housing companies is reported. the energy profile indicators and the metered energy consumption values were collected for 155 building entities (mostly 3the type of energy calculation procedure has only a very small effect (for example the seasonal method has only an uncertainty of ±2 % compared to the monthly method [14]). thus, nearly the same uncertainties would result if a highresolution simulation programme was used. large building blocks) with 3329 apartments4. the energy characteristics of the building sample comprise a wide span from unrefurbished to deep renovations [12], which also includes the use of passive house components and installation of ventilation systems with heat recovery. in two of the three companies, a large effort was required to collect the information from individual building-related files. in the third company, however, energy profile indicators were already introduced in 2008 for the complete portfolio and since then maintained and updated [3]. these data could therefore be used directly. for some of the buildings, the indicators were collected during a quick on-site inspection – the typical effort was half an hour per building. table 4 shows the results for a subset where metered consumption is available for both, heating and dhw. as a parameter for the theoretical heat loss the “thermal conductance” is defined, which is the heat loss of the building per kelvin temperature difference, related to the reference floor area. it can be stated that the average actual consumption by conductance class is very close to the calculation model averages (the ratio average actual to average theoretical energy use varies from 95 % to 115 % per class). the standard deviations of the actual energy consumption and the average uncertainty of calculation are in similar ranges. in summary, the method seems quite suitable for modelling the consumption of housing stocks and for evaluating measures of different depths with respect to their actual effect. one conclusion that can be drawn from this is, for example, that the average energy consumption for heating and dhw of the class with best insulation level is about one third of that with the worst. 6. conclusions a method has been developed that is based on a set of monitoring indicators collectable by a questionnaire or a quick on-site examination. these “energy profile indicators” are used as input for an energy performance calculation to determine an expected range of actual energy consumption. examples show how the uncertainty of consumption prognosis depends on information about data source and data completeness. moreover, the methodology was applied to a sample of multi-family buildings involving statistical evaluations. for different levels of insulation, the estimated as well as the actually measured consumption were determined, with the actual consumption by conductance class being very close to the calculation model averages. future work should focus on the expansion of the database to include also single-family houses and different heat supply systems, especially heat pumps and solar systems. 4the datasets of the sample can be viewed at the excel workbook mentioned above [8]. 475 t. loga, b. stein, g. behem acta polytechnica ctu proceedings thermal conductance class∗ range w/(m2k) 0.01–0.80 0.81–1.00 1.01–1.20 1.21–1.50 1.51–2.00 2.01– equivalent insulation thickness of opaque elements∗∗ average cm 31.6 21.5 12.1 13.0 2.4 2.7 u-value of windows average w/(m2k) 0.80 1.29 1.52 2.62 2.68 2.68 availability of mechanical ventilation with heat recovery percentage 100 % 0 % 0 % 0 % 0 % 0 % calculated energy use for heating and dhw (comparison value)∗∗∗ average kwh/(m2a) 50 73 96 103 143 163 std. dev. kwh/(m2a) ±7 ±5 ±10 ±11 ±14 ±20 av. uncert. kwh/(m2a) ±16 ±20 ±22 ±27 ±39 ±43 actual energy use for heating and dhw (metering) average kwh/(m2a) 50 78 110 110 135 148 std. dev. kwh/(m2a) ±9 ±21 ±15 ±21 ±16 ±23 ratio actual / calculated energy use 0.99 1.07 1.15 1.07 0.95 0.91 frequency number of building n = 15 n = 10 n = 8 n = 23 n = 5 n = 11datasets number of annual n = 28 n = 25 n = 21 n = 46 n = 10 n = 22consumption values ∗ the “thermal conductance” is the heat transfer coefficient by transmission and ventilation, related to the reference floor area of the building. ∗∗ thermal resistance of opaque envelope, expressed in cm insulation with 0.035 w/(m · k) thermal conductivity. ∗∗∗ energy use, calculated by the combined probability-physics model, balance scope and climate consistent with metering (example: if heat is metered in the apartments the heat losses of the central heat distribution system are disregarded in this comparison). table 4. comparison of theoretical and actual consumption for a building sample (heating + dhw). averages, uncertainties and standard deviations differentiated by thermal conductance class. (status of database analysis: 2022-02-17). in practice, the simplification of data input, the ability to adequately consider data gaps and uncertainties and the output of estimation ranges may provide benefits for different purposes: • individual houses (occupants, owners, energy consultants, . . . ): ▷ the small effort to provide the indicators for an existing house makes it a useful method in energy advice campaigns. the current energy consumption can be interpreted and assessed. achievable consumption bandwidths after refurbishment can be provided for the investigated house. ▷ for refurbished buildings as well as for new build, the method provides a target range of the energy consumption. the occupant or owner can compare the actual consumption against this individual benchmark. if the consumption is above the expectation range, the implemented measures, the operating conditions, and the occupant behaviour can be checked and improved, if necessary. besides, measurements of the actual utilisation conditions and of the local climate during the consumption period can help to significantly narrow down the expectation range. ▷ energy profile indicators can be attached to epc issuing (e.g. “statistic forms”, xml-files). this enables a plausibility control of the energy performance calculated for the epc, for issuers as well as for inspectors on quality assurance. ▷ the fact that visually provable indicators are used is also a benefit for application cases in the context of pricing rules, dependent on energy performance (building value, financing, sales, rents, . . . ) and for grant applications. • housing portfolios (housing companies, energy service and billing companies, . . . ): ▷ the collection and regular update of energy performance indicators provides a basis to determine the current level of insulation and the current fraction of efficient/renewable systems for the entire stock as well as the respective annual improvement rates. these achievements can e.g. be compared with the climate protection targets for the portfolio or for the national stock. ▷ the average estimated and, if available, actual consumption values and ranges per conductance class are benchmarks that can be useful in the preparation of refurbishments, especially in the communication with tenants, to strengthen the confidence in the actual effect of the measures. ▷ by matching consumption values of single buildings against the benchmarks, cases with suspiciously high consumption can be identified. an uptake of low-cost measures can improve the situation, similar to the target/actual comparison described for individual buildings. ▷ the calculation for all buildings of a portfolio may also be used as building stock model for developing refurbishment strategies. • national housing stock (energy experts, authorities, energy agencies, . . . ): ▷ by using the energy profile indicators for a random-sample-based house owner survey, the achievements regarding levels of insulation and fractions of efficient/renewable systems can be determined for the national housing stock, similar to the evaluation of the energy profile database of a housing portfolio. 476 vol. 38/2022 use of energy profile indicators to determine the expected range . . . ▷ in addition, the random sample can be used in national building stock models in form of a microsimulation model or by aggregation to average buildings to forecast the development of energy use by energy carrier and calculate different alternative scenarios. ▷ in the context of national information campaigns, the publication of benchmarks consisting of average consumption values and typical spreads for different insulation categories can provide useful information for occupants and owners that inspires confidence in different types of measures. acknowledgements the work described in this paper is part of the research project mobasy, supported by the german government (federal ministry for economic affairs and climate action (reg. n°: 03sbe0004a); project duration: 11/2017–10/2022; more information at: https: //www.iwu.de/forschung/energie/2017/mobasy/). the analysis of the housing sample has been implemented in collaboration with three german housing companies: bauverein ag darmstadt, wohnbau gießen and nassauische heimstätte wohnstadt frankfurt/main. we very much appreciate the support and cooperation. references [1] t. loga, m. großklos, a. müller, et al. realbilanzierung für den verbrauch-bedarf-vergleich (mobasy teilbericht). tech. rep., darmstadt, 2021. isbn 978-3-941140-67-7. https://doi.org/10.13140/rg.2.2.22472.24328/1 [2] t. loga, n. diefenbach, j. knissel, r. born. entwicklung eines vereinfachten, statistisch abgesicherten verfahrens zur erhebung von gebäudedaten für die erstellung des energieprofils von gebäuden („kurzverfahren energieprofil“). tech. rep., darmstadt, 2005. isbn 3-932074-76-9. https://doi.org/10.13140/rg.2.2.20794.52160 [3] t. loga, m. repp, s. swiderek. verbrauchsbenchmarks – ein soll-ist-vergleich für wohnungsunternehmen. in 12. effizienztagung bauen + modernisieren 2020. 2020. [2022-03-02]. https://www.researchgate.net/ publication/350710240_verbrauchsbenchmarks__ein_soll-ist-vergleich_fur_wohnungsunternehmen [4] t. loga, k. müller. evaluation of the tabula database – comparison of typical buildings and heat supply systems from 12 european countries, iwu, darmstadt, 2012. tabula work report, [2022-03-02]. http://www.buildingtypology.eu/downloads/public/docs/report/ tabula_workreport_evaluationdatabase.pdf [5] n. diefenbach, h. cischinsky, m. rodenfels, k.-d. clausnitzer. datenbasis gebäudebestand. datenerhebung zur energetischen qualität und zu den modernisierungstrends im deutschen wohngebäudebestand, iwu, darmstadt, 2010. [2022-10-06]. https://www.iwu.de/fileadmin/ publikationen/gebaeudebestand/datenbasis/ endbericht_datenbasis.pdf [6] h. cischinsky, n. diefenbach. datenerhebung wohngebäudebestand 2016, iwu, darmstadt, 2018. [2022-10-06]. https://www.iwu.de/fileadmin/ publikationen/gebaeudebestand/2018_iwu_ cischinskyetdiefenbach_datenerhebungwohngeb%c3%a4udebestand-2016.pdf [7] t. loga. energy profile. questionnaires thermal envelope and heat supply system, iwu, darmstadt, 2021. [2022-03-02]. completed with example data: https: //www.iwu.de/fileadmin/tools/energyprofile/ energyprofile_forms_buildingsystem_eng.pdf template with active form fields: https://www.iwu.de/ fileadmin/tools/energyprofile/energyprofile_ formbuildingsystem_datacapture.pdf [8] t. loga. energy profile indicators. data table, showcase, iwu, darmstadt, 2021. [2022-03-02]. https://www.iwu.de/fileadmin/tools/ energyprofile/energyprofileindicators_ datatable_showcase_eng.xlsx [9] w. tian, y. heo, p. de wilde, et al. a review of uncertainty analysis in building energy assessment. renewable and sustainable energy reviews 93:285–301, 2018. https://doi.org/10.1016/j.rser.2018.05.029 [10] t. loga, b. stein, g. behem. cesb22 – annex of the paper: use of energy profile indicators to determine the expected range of heating energy consumption, cesb22 prague, documentation of the example building calculations, 2022. [2022-03-02]. https://www.iwu.de/fileadmin/publikationen/ energie/mobasy/cesb22_energyprofileindicators_ annex_showcaseexamples_calculationflow.pdf [11] t. loga, n. diefenbach, j. knissel. entwicklung eines flächenschätzverfahrens zur abschätzung von bauteilflächen als grundlage für die energetische bewertung von wohngebäuden. bauphysik 27(2):102– 116, 2005. https://doi.org/10.1002/bapi.200590017 [12] t. loga, g. behem. target/actual comparison and benchmarking used to safeguard low energy consumption in refurbished housing stocks. in proceedings of the eceee summer study 2021, pp. 461–471. 2021. [2022-03-02]. https://www.researchgate.net/ publication/355124720_targetactual_comparison_ and_benchmarking_used_to_safeguard_low_energy_ consumption_in_refurbished_housing_stocks [13] t. loga, n. diefenbach. tabula calculation method – energy use for heating and domestic hot water. reference calculation and adaptation to the typical level of measured consumption, iwu, darmstadt, 2013. tabula documentation, [2022-03-02]. https://episcope.eu/fileadmin/tabula/public/ docs/report/tabula_commoncalculationmethod.pdf [14] t. loga. heizgrenztemperaturen für gebäude unterschiedlicher energetischer standards. abschätzung der heizgrenztemperaturen und bilanzzeiten für das heizperiodenbilanzverfahren nach en 832/din v 4108-6. bauphysik 25(1):2–14, 2003. https://doi.org/10.1002/bapi.200300020 477 https://www.iwu.de/forschung/energie/2017/mobasy/ https://www.iwu.de/forschung/energie/2017/mobasy/ https://doi.org/10.13140/rg.2.2.22472.24328/1 https://doi.org/10.13140/rg.2.2.20794.52160 https://www.researchgate.net/publication/350710240_verbrauchsbenchmarks_-_ein_soll-ist-vergleich_fur_wohnungsunternehmen https://www.researchgate.net/publication/350710240_verbrauchsbenchmarks_-_ein_soll-ist-vergleich_fur_wohnungsunternehmen https://www.researchgate.net/publication/350710240_verbrauchsbenchmarks_-_ein_soll-ist-vergleich_fur_wohnungsunternehmen http://www.building-typology.eu/downloads/public/docs/report/tabula_workreport_evaluationdatabase.pdf http://www.building-typology.eu/downloads/public/docs/report/tabula_workreport_evaluationdatabase.pdf http://www.building-typology.eu/downloads/public/docs/report/tabula_workreport_evaluationdatabase.pdf https://www.iwu.de/fileadmin/publikationen/gebaeudebestand/datenbasis/endbericht_datenbasis.pdf https://www.iwu.de/fileadmin/publikationen/gebaeudebestand/datenbasis/endbericht_datenbasis.pdf https://www.iwu.de/fileadmin/publikationen/gebaeudebestand/datenbasis/endbericht_datenbasis.pdf 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https://www.iwu.de/fileadmin/tools/energyprofile/energyprofile_formbuildingsystem_datacapture.pdf https://www.iwu.de/fileadmin/tools/energyprofile/energyprofile_formbuildingsystem_datacapture.pdf https://www.iwu.de/fileadmin/tools/energyprofile/energyprofileindicators_datatable_showcase_eng.xlsx https://www.iwu.de/fileadmin/tools/energyprofile/energyprofileindicators_datatable_showcase_eng.xlsx https://www.iwu.de/fileadmin/tools/energyprofile/energyprofileindicators_datatable_showcase_eng.xlsx https://doi.org/10.1016/j.rser.2018.05.029 https://www.iwu.de/fileadmin/publikationen/energie/mobasy/cesb22_energyprofileindicators_annex_showcaseexamples_calculationflow.pdf https://www.iwu.de/fileadmin/publikationen/energie/mobasy/cesb22_energyprofileindicators_annex_showcaseexamples_calculationflow.pdf https://www.iwu.de/fileadmin/publikationen/energie/mobasy/cesb22_energyprofileindicators_annex_showcaseexamples_calculationflow.pdf https://doi.org/10.1002/bapi.200590017 https://www.researchgate.net/publication/355124720_targetactual_comparison_and_benchmarking_used_to_safeguard_low_energy_consumption_in_refurbished_housing_stocks https://www.researchgate.net/publication/355124720_targetactual_comparison_and_benchmarking_used_to_safeguard_low_energy_consumption_in_refurbished_housing_stocks https://www.researchgate.net/publication/355124720_targetactual_comparison_and_benchmarking_used_to_safeguard_low_energy_consumption_in_refurbished_housing_stocks https://www.researchgate.net/publication/355124720_targetactual_comparison_and_benchmarking_used_to_safeguard_low_energy_consumption_in_refurbished_housing_stocks https://episcope.eu/fileadmin/tabula/public/docs/report/tabula_commoncalculationmethod.pdf https://episcope.eu/fileadmin/tabula/public/docs/report/tabula_commoncalculationmethod.pdf https://doi.org/10.1002/bapi.200300020 acta polytechnica ctu proceedings 38:470–477, 2022 1 introduction 2 energy profile indicators 3 simple combined physics-probability model 3.1 principle 3.2 systematic uncertainty assessment 3.3 mechanisms of the energy performance calculation based on energy profile indicators 4 showcase: influence of different parameters on the uncertainty of the estimated consumption 5 application on a sample of apartment blocks from housing companies 6 conclusions acknowledgements references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.37.0048 acta polytechnica ctu proceedings 37:48–53, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague on the ongoing computation of the lvr-15 reactor fuel burnup jan pintaa, b a czech technical university in prague, faculty of nuclear sciences and physical engineering, department of nuclear reactors, v holešovičkách 2, 180 00 prague 8, czech republic b research centre řež, department of neutron physics, hlavní 130, 250 68 husinec-řež, czech republic correspondence: pintajan@fjfi.cvut.cz abstract. isotopic composition of the lvr-15 reactor fuel changes significantly during the operation time. in order to enable monte carlo calculations of the lvr-15 core throughout any reasonable time period, the isotopic compositions of partially depleted fuel assemblies have to be known at the start of the simulation. even though the infinite lattice calculation can represent useful initial approximation, the non-leakage assumption is not necessarily valid for compact and highly heterogenous cores like in case of the lvr-15. therefore, a more advanced calculation process have to be implemented towards appropriate determination of fuel assembly burnup. the main aim of this paper is to describe the current state of the ongoing computation as well as provide an explanation of the chosen calculation scheme of the lvr-15 operation history. currently available results that are affected by profound inaccuracy are discussed in the final part alongside with planned future improvements. keywords: burnup calculation, serpent, lvr-15 reactor. 1. introduction the lvr-15 reactor [1] is a tank-type light water research reactor with 10 mw thermal power output. the reactor utilizes the irt-4m 19.7 % enriched fuel and its main applications are material testing, silicon doping and production of radioisotopes. currently, the serpent [2] computational model of the reactor is being developed with aim to perform its experimental qualification in the near future. one of the most crucial challenges concerning the development is a coverage of fuel burnup. there is an essential question of how to create proper representations of partially depleted fuel assemblies that are present at the start of an operation cycle. in fact, there is a majority of fuel assemblies with nonzero burnup contained inside the lvr-15 reactor core at the beginning of each cycle. therefore, determination of isotopic composition of partially depleted fuel have to be implemented in order to obtain convincing initial conditions that are needed for the serpent calculation of any time step of any lvr-15 reactor cycle. the first obvious possibility is to calculate burnup-wise fuel isotopic concentrations in the infinite lattice approximation. however, it will be shown later that such approach does not have to be necessarily valid for compact and highly heterogenous cores with considerable neutron leakage. for this reason, a more advanced fuel burnup computation procedure has been applied. the idea of this procedure is to take into account two important characteristics of the lvr-15 reactor. the first one is the above mentioned compactness and high heterogeneity of the core. the second one is the complicated operation scheme of the reactor, which involves short-term shutdowns in the run of a cycle operation. by adopting these two features, a more exact approximation in the fuel burnup implementation should be obtained. this work is particularly focused on the introduction and description of the advanced lvr-15 reactor fuel burnup computation. the results presented in this paper are considered rather preliminary due to the fact that the computational process is still in the ongoing state. nevertheless, the profound improvements will be realized on the basis of these first results in the following iteration of the fuel burnup calculation. 2. serpent model the current version of the 3d serpent model of the lvr-15 reactor allows to build the reactor model with any desired core configuration that contains the most frequently used components, i.e. fuel assemblies, beryllium blocks, water displacers and irradiation channels. all components in the core are surrounded with water and placed inside the reactor vessel. an instance of a core configuration geometry plot of the serpent model is provided in figure 1. as for fuel components, both standard 8-tube and control 6-tube irt-4m fuel assemblies can be chosen. control rod positions are adjustable according to the demanded pattern. in addition, both fuel assemblies can be defined with a variety of possible burnup distribution structures. different fuel material compositions can be assigned to corresponding regions of a particular fuel component using these distribution structures. 48 https://doi.org/10.14311/app.2022.37.0048 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 37/2022 on the ongoing computation of the lvr-15 reactor fuel burnup figure 1. serpent geometry plot of the opening lvr-15 core configuration of the k216 cycle; two fuel assemblies with initial zero burnup are highlighted (standard 85a710 – red, control 65c666 – black). the initial fuel burnup implementation is based on burnup-wise isotopic compositions of standard fuel assembly calculated with serpent in axially and radially infinite lattice. the serpent model utilizes predefined burnup library with atomic densities of considered nuclides as functions of burnup (see table 1). 3. computational process the chosen process of the burnup computation is based on the real operation scheme of the lvr-15 research reactor. the complexity of the reactor schedule is emphasized in figure 2 and figure 3, where powers and keff from serpent are visualized, respectively. more than a 2.5 year latest time period of the reactor operation has been considered for the purpose of the computation. this time interval was determined according to the moment when a fuel assembly with the largest burnup in the currently running cycle has been introduced into the reactor core for the first time as a fresh fuel component. this condition is fulfilled by the k216 cycle, where two new fuel assemblies have been added into the operation procedure – one standard (85a710) and one control (65c666). the position of the former is marked with red colour, while the latter is highlighted with black colour in terms of the k216 core layout visualization in figure 1. the whole computational process consists a total of 14 operation cycles starting from the k216 up to the k229. two fresh fuel assemblies in average are introduced into the reactor core at the start of each of the examined cycles. since that point, burnup development of these fuel assemblies can be calculated without the uncertainty in the fuel material composition. this presumption is supported by the fact that the fresh fuel assemblies have known initial isotopic concentrations. overview of fuel assemblies with this attribute is listed in table 2 alongside with the information whether any particular fuel assembly was present in the corresponding operation cycle. the computational process of the latest operation history of the lvr-15 reactor has been designed on the basis of data from the nodal code that is used for cycle designing at the lvr-15 [3]. these data come from nodal calculation of given cycle that has been carried out with genuine operational power data of the reactor. a typical calculation schedule of the lvr-15 reactor cycle contains over a hundred partial critical calculation steps. that reflects the complexity of the reactor operation. the same approach has been adopted in the serpent based computational process in order to achieve authentic operational conditions of the lvr-15 reactor as closely as possible. as for initial step of the entire computation, referred as the step 0 of the k216 cycle, the material compositions of fuel assemblies with nonzero burnup have been obtained from the above mentioned infinite lattice burnup library implemented in the serpent computational model. axial-wise burnup data divided into 5 regions of these fuel assemblies have been taken from the calculation of the previous k215 cycle from the operational nodal code. the scheme of the computational process is expressed in figure 4. a single step of the computation other than the initial step 0 of the k216 cycle can be described with a sequence of the same partial procedures. first, the appropriate operational data of the lvr-15 are loaded and transformed into the input file that is needed for the building of the corresponding lvr-15 reactor serpent model. these data include: • the name of the current cycle, • the calculation step number in the cycle, • time interval length of the step, • reactor power during the calculation step, • identifiers of present fuel assemblies, • core configuration layout within the step, • control rod insertions in the step core layout, • specifics of uranium targets in the core centre. in addition to the input file, the second set of data has to be provided. the second set contains calculated isotopic compositions of fuel assemblies from the previous step. this data stream is referred as the olddepfile and it is sent to the slvr15 together with the input file. the slvr15 procedure represents a software tool designed for creation of the serpent input file of demanded configuration of the lvr-15 reactor 49 jan pinta acta polytechnica ctu proceedings 80160 130270 420950 430990 441010 451030 471090 481130 531350 541350 551330 551340 551350 601430 601450 611470 611490 611510 621470 621490 621500 621510 621520 631530 631550 641550 641570 922340 922350 922360 922380 932370 942380 942390 942400 942410 942420 952410 952430 962430 962440 962450 table 1. list of 42 considered nuclides that are passed on between subsequent computation steps (zai numbers [4]). k216 k217 k218 k219 k220 k221 k222 k223 k224 k225 k226 k227 k228 k229 85a710 × × × × × × × × – × × × × × 65c666 × × × × × × × × × × × × × × 85a711 – × × × × × × × × × × × × × 85a712 – × × × × × × × – × × × × × 65c667 – × × × × × × × × × × × × × 65c668 – – × × × × × × × × × × × × 85a713 – – – × × × × × × × × × × × 85a714 – – – × × × × × × × × × × × 65c669 – – – – × × × × × × × × × × 85a715 – – – – – × × × × × × × × × 65c670 – – – – – × × × × × × × × × 65c671 – – – – – – × × × × × × × × 65c672 – – – – – – × × × × × × × × 85a716 – – – – – – – × × × × × × × 65c673 – – – – – – – × × × × × × × 85a717 – – – – – – – – × × × × × × 85a718 – – – – – – – – × × × × × × 65c674 – – – – – – – – × × × × × × 65c675 – – – – – – – – – × × × × × 85a719 – – – – – – – – – – × × × × 85a720 – – – – – – – – – – × × × × 85a721 – – – – – – – – – – – × × × 65c676 – – – – – – – – – – – × × × 60c834 – – – – – – – – – – – – × × 60c835 – – – – – – – – – – – – × × 85a722 – – – – – – – – – – – – – × 85a723 – – – – – – – – – – – – – × table 2. the presence of initially fresh fuel assemblies throughout considered cycles k216–k229 (× in the cycle, – not in the cycle). 0 5 10 k216 k217 k218 k219 k220 k221 k222 0 14 28 0 5 10 k223 14 28 k224 14 28 k225 14 28 k226 14 28 k227 14 28 k228 14 28 k229 time (days) po w er (m w th ) figure 2. power sequence of the lvr-15 reactor operation including cycles k216–k229. 50 vol. 37/2022 on the ongoing computation of the lvr-15 reactor fuel burnup 0.9 1.0 k216 k217 k218 k219 k220 k221 k222 0 14 28 0.9 1.0 k223 14 28 k224 14 28 k225 14 28 k226 14 28 k227 14 28 k228 14 28 k229 time (days) k e ff figure 3. development of keff calculated with serpent throughout cycles k216–k229. lv r -1 5 op er at io n da ta input slvr15 olddepfile mainserpent newdepfile step n step n+1 figure 4. scheme of the computational process. computational model. the main input file is prepared for the serpent calculation of particular step of the computational process after the slvr15 execution. one of the step calculation outputs is the serpent depletion file indicated as the newdepfile. this output file alongside with appropriate operational data is used as the input for the calculation of the following step. the introduced process is repeated n times according to the number of steps in the whole computation. the isotopic composition of fuel assemblies between subsequent steps is passed on for 42 selected nuclides. the list of considered nuclides is written in table 1. the decision of the usage of the 42 nuclide approximation was made based on the burnup credit isotopes [5] with consideration of 135xe. the simulation of the xenon poisoning problematics was achieved by the elimination of 135i and 135xe nuclei before the beginning of each cycle, i.e. k216, k217, k218, etc. this behaviour was assumed to be sufficient enough to simulate actual operational conditions in the lvr-15 reactor core. parameters for neutron population in criticality source mode of each individual serpent calculation were set to 100 000 neutrons per generation, 200 active and 40 inactive generation cycles. standard deviation of keff did not exceed 27 pcm for each calculation step. all serpent calculations were carried out using the endf/b-viii.0 [6] based cross section library. 4. results the collection of three entangled types of results is presented in this paper. the first type of results covers calculated atomic densities of chosen nuclide compared to the infinite lattice approximation. second type of results represents verification of the calculated atomic densities with aim to determine whether the results are credible or not. the final part of the collection supports the statement that has been gained from the previous type of the results. 4.1. calculated burnup qualification visualization of 239pu atomic density burnup dependence has been chosen for the purpose of the calculated fuel burnup qualification. the axially averaged 239pu concentrations in the fuel as functions of burnup are provided in figure 5. these results are taken from the advanced computational process described in section 3. the visualization of the data is shown for all fuel assemblies with initial zero burnup at the moment of their first introduction into the lvr-15 reactor core during studied cycles k216–k229. the comparison with the infinite lattice calculated 239pu atomic densities is also included in figure 5. 51 jan pinta acta polytechnica ctu proceedings 0 20 40 60 80 100 burnup (mwd/kgu) 0.0e+00 1.1e-06 2.3e-06 3.4e-06 4.5e-06 at om ic d en si ty (1 e2 4/ cm 3) 0.0e+00 1.1e-05 2.3e-05 3.4e-05 4.5e-05 in fin ite la tt ic e at om ic d en si ty (1 e2 4/ cm 3) infinite lattice 85a711 65c666 85a710 65c667 85a712 85a713 85a714 65c668 85a715 65c669 65c672 65c671 65c670 85a716 85a717 65c673 65c675 65c674 85a718 85a719 85a720 65c676 85a721 60c835 60c834 figure 5. axially averaged atomic densities of 239pu calculated for all fuel assemblies with initial zero burnup (left axis). the calculated values are compared with atomic densities from the infinite lattice approximation (right axis). 0 50 100 150 1.1e-04 7.5e-04 1.4e-03 u-235 0 50 100 150 5.3e-03 5.4e-03 5.5e-03 u-238 0 50 100 150 0.0e+00 2.3e-05 4.5e-05 pu-239 0 50 100 150 0.0e+00 3.8e-06 7.6e-06 pu-241 burnup (mwd/kgu) at om ic d en si ty (1 e2 4/ cm 3) infinite lattice 85a710 65c666 figure 6. comparison of assembly-wise atomic densities between the infinite lattice approximation and the unbroken burnup simulation of the opening step of the k216 cycle. two fuel assemblies, standard 85a710 and control 65c666, has been taken into the account in the latter type of the calculation. values for 235u, 238u, 239pu and 241pu are included. a considerably profound discrepancies between the calculated results and the infinite lattice approximation can be observed in figure 5. it is apparent that the amount of 239pu in case of the infinite lattice is one order of magnitude higher than in the results from the lvr-15 reactor operation history computation. the result of such kind raises the immediate demand for a validation of the entire computational process and for a verification of the data as well. the attempt to clarify these incompatibilities is the point of interest of the following type of the results collection. 4.2. computational process validation the validation of the actual computational process has been carried out based on the previous results. therefore, two kinds of problems have been compared. first, the burnup calculation of the infinite lattice approximation that was used as a prerequisite for the considered burnup library mentioned in section 2. second, an uninterrupted burnup calculation of the lvr-15 reactor core in the initial point of the whole computational process, i.e. the step 0 of the k216 cycle configuration. this comparison also represents the difference in the non-leakage approach against the genuine core configuration calculation with emphasis on high heterogeneity and compactness of the lvr-15 core. the data from the unbroken calculation of the opening step of the k216 cycle involves axially averaged atomic densities of the fresh standard and control fuel assemblies with identifiers 85a710 and 65c666, respectively. altogether, atomic density plots of 4 nuclides, 235u, 238u, 239pu and 241pu, are illustrated in figure 6. the greatest similarity of the atomic density data is observable in case of 235u. still, figure 6 shows that the depletion of 235u is slightly more significant in case of both fuel assemblies from the unbroken cycle simulation. however, the differences for 238u and for both pictured isotopes of plutonium are more extensive. the production of plutonium is directly linked to the neutron capture on 238u nuclei. the figure 6 indicates that there is a noticeably higher amount of 238u capture in the infinite lattice approximation and therefore a lesser concentration of the plutonium isotopes in the fuel assemblies of the comparison k216 cycle calculation. to proceed to the main point of this validation analysis, there is a fundamental assumption that the 239pu atomic densities should be in the same order of magnitude as the atomic densities determined from the unbroken real core burnup calculation. such accordance, however, has not been achieved. this outcome gives an evidence that some presumption in the designed computational process has been applied inaccurately. the investigation of the cause of the expressed discrepancies in the computational process is carried out in the following type of the results collection. 4.3. the cause of discrepancies the transfer of fuel composition data between particular steps of the computational process has been investigated as a potentially problematic part in the process design. more accurately, in the determination of which nuclides should be chosen for the conservation of atomic densities from one step to another. originally, there were 42 chosen nuclides as stated in table 1. a new calculation concerning a first few steps of the carried out computational process was executed with the change in the number of chosen nuclides. this time, atomic densities of all nuclides in the serpent cross section library from the end of a particular calculation step have been written into the input data of the following step. due to this decision, the expected results have changed significantly 52 vol. 37/2022 on the ongoing computation of the lvr-15 reactor fuel burnup 0.0 0.2 0.4 0.6 0.8 1.0 burnup (mwd/kgu) 0.0e+00 6.5e-08 1.3e-07 2.0e-07 2.6e-07 3.3e-07 3.9e-07 at om ic d en si ty (1 e2 4/ cm 3) infinite lattice k216 unbroken 42 nuclides all nuclides figure 7. short-term development of 239pu atomic densities for various calculation types. towards the results that were observed in figure 6. this behaviour for the case of 239pu atomic densities is shown in figure 7. 5. discussion according to the presented results, a misleading assumption in the design of the advanced computational process has been made. it has been investigated that the 42 nuclide representation does not reflect realistic conditions in the core during the lvr-15 reactor operation and thus a more comprehensive attitude has to be implemented. clearly, the consideration of all nuclides available in the serpent cross section library should be involved in the next generation of the lvr-15 reactor burnup computation. even though such upgrade of the computational process will require a larger amount of space for the output data storage, running time of the serpent calculation of a single step will not be affected. this indicates the feasibility of the improvement. on the other hand, the obtained results in figure 5 shows the differences in the fuel composition based on the individual operation scheme of each fuel assembly. therefore, the data can be used for analysis of the influence of the various schemes on the final form of the atomic density that will be represented in the burnup library of the lvr-15 reactor serpent model. 6. conclusions the serpent computational model of the lvr-15 reactor has been introduced in this work with focus on the problematics of fuel burnup implementation. the advanced burnup computational process has been described using the most important aspects and operational characteristics of the lvr-15 reactor. currently available results have been presented and analyzed according to the additional calculations. the analysis has shown that the number of nuclides considered in the transfers between calculation steps have been chosen inaccurately. this misleading nuclide interpretation has caused significant discrepancies in the atomic density values between the computation and the fundamentally anticipated results from the infinite lattice approximation and from the unbroken burnup calculation of the realistic lvr-15 core. the outcome of this paper will be used in the next generation of the advanced burnup computational process in order to determine a more realistic burnupwise fuel composition of the lvr-15 reactor. list of symbols keff effective multiplication factor endf evaluated nuclear data file references [1] m. koleška, z. lahodová, j. šoltés, et al. capabilities of the lvr-15 research reactor for production of medical and industrial radioisotopes. journal of radioanalytical and nuclear chemistry 305:51–59, 2015. https://doi.org/10.1007/s10967-015-4025-5. [2] j. leppänen, m. pusa, t. viitanen, et al. the serpent monte carlo code: status, development and applications in 2013. annals of nuclear energy 82:142–150, 2015. https://doi.org/10.1016/j.anucene.2014.08.024. [3] j. ernest. abstract of computational program noder 2006. [4] j. leppänen, et al. serpent – a continuous-energy monte carlo reactor physics burnup calculation code. vtt technical research centre of finland 4, 2013. [5] nea, oecd. spent nuclear fuel assay data for isotopic validation 2011. state-of-the-art report, https://www.oecd-nea.org/science/wpncs/adsnf/ soar_final.pdf. [6] d. brown, m. chadwick, r. capote, et al. endf/b-viii.0: the 8th major release of the nuclear reaction data library with cielo-project cross sections, new standards and thermal scattering data. nuclear data sheets 148:1–142, 2018. https://doi.org/10.1016/j.nds.2018.02.001. 53 https://doi.org/10.1007/s10967-015-4025-5 https://doi.org/10.1016/j.anucene.2014.08.024 https://www.oecd-nea.org/science/wpncs/adsnf/soar_final.pdf https://www.oecd-nea.org/science/wpncs/adsnf/soar_final.pdf https://doi.org/10.1016/j.nds.2018.02.001 acta polytechnica ctu proceedings 37:48–53, 2022 1 introduction 2 serpent model 3 computational process 4 results 4.1 calculated burnup qualification 4.2 computational process validation 4.3 the cause of discrepancies 5 discussion 6 conclusions list of symbols references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.37.0001 acta polytechnica ctu proceedings 37:1–9, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague subsals: a subchannel thermal-hydraulic code for irt type fuel analysis tomáš adámeka, b a czech technical university in prague, faculty of nuclear sciences an physical engineering, department of nuclear reactors, břehová 7, 115 19 prague, czech republic b research centre řež, hlavní 130, husinec-řež, czech republic correspondence: tomas.adamek@cvrez.cz abstract. the lvr-15 research reactor is operated with a tube type fuel irt-4m. due to fuel’s unique concentric square annular shape with, coolant flow is subject to significant pressure driven crossflow. detailed calculation of such flow patterns is beyond capabilities of standard system codes used for the thermal hydraulic safety analysis. to assess safety of core designs consisting of the irt-4m fuel assemblies, a new subchannel code is under development at research centre řež and czech technical university in prague. computer program subsals is single phase, steady state, subchannel solver in development. rigid computational mesh describes irt-like geometries, and it is also suitable for other specific doubly connected ducts. in this contribution, general description of calculation procedure and computational capabilities is presented along with code-to-code comparison. it could be used as proof of concept for future improvements in thermal hydraulic calculations of research reactors. keywords: subsals, subchannel analysis, thermal-hydraulics, irt-4m, lvr-15, crossflow. 1. introduction thermal-hydraulic analysis of nuclear reactors is a necessary part of safety assessment. to maintain the integrity of physical barriers during the operation, appropriate cooling of all fuel assemblies must be ensured. various sophisticated computational tools have been developed for the simulation of water cooled reactors behaviour. over the past decades increasing computational power of the computers allowed for a great development of one-dimensional system codes. nuclear reactor core thermal-hydraulic safety assessment based on a hot channel theory, adopted to use in system codes, may be unnecessarily conservative. a more realistic approach to evaluate a core performance is accomplished using the subchannel analysis techniques. computational models can be used to represent either an entire core or an individual sub-assembly. the subchannel approach has been used extensively in light water reactor applications. it provides a higher degree of resolution than traditional system code approach and allows for a fast calculation of the flow, temperature, and void distributions with a reasonable accuracy. main domain of application of the traditional subchannel codes like subchanflow are core designs consisting of the rod bundles. moreover, the efforts are made to develop advanced computational tools for the thermal-hydraulic analysis of the reactors operated with plate type fuel. for the utilization in reactors using tube type fuel, especially with significant pressure driven crossflow, however, modelling capabilities of the standard thermal-hydraulic simulation tools are limited. 2. general description of the subchannel solver subsals computer program subsals (subchannel square annular layout solver) is a single phase, steady state subchannel solver under development at research centre řež and czech technical university in prague. its main purpose is the thermal-hydraulic analysis of the lvr-15 research reactor, which main goal is to determine cladding temperature distribution in the irt-4m fuel assemblies under the normal operation. 2.1. irt type fuel and computational mesh the irt type fuel is a series of sandwich, tube-type fuel assemblies with a square cross section, originally designed in 1963 as the irt-m consisting of three fuel tubes. later, further improvements in the fabrication process were achieved which led to four tube design irt-2m. in 1979 a new generation, the irt-3m was developed with an enhanced heat transfer area. finally in 2004, the irt-4m (figure 1) design motivated by the rertr program was completed [1]. currently the irt type fuel is used in two research reactors in czech republic (lvr-15 and vr-1), two research reactors in russia (ir-8 and irt-t) and one research reactor in uzbekistan (vvr-sm) [2]. 1 https://doi.org/10.14311/app.2022.37.0001 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en tomáš adámek acta polytechnica ctu proceedings figure 1. geometry of the irt-4m fuel assemblies: standard 8-tube fuel assembly (top) and control 6-tube fuel assembly (bottom) [3]. 0 1 2 3 4 5 6 7 8 0 1 2 3 4 5 6 7 8 1 20 15 23 23 22 22 14 21 21 5 4 3 6 9 9 10 10 7 8 12 11 11 12 13 131415 1415 91011 17 17 16 16 18 12 18 19 19 20 20 13 figure 2. radially-azimuthal view of the computational mesh. in it’s current state of development, for the sake of modularity, subsals is designed to analyse one coolant channel between the fuel tubes. the interaction between multiple coolant channels will be achieved by the external iterator module in later stages of development. the computational mesh (figure 2) is rigid in radial and azimuthal direction and user defined in axial direction. radially one coolant channel azimuthally divided into 16 unique subchannels is presented surrounded from both sides by fuel tubes, both azimuthally divided into 24 unique fuel elements. 2.2. basic conservation equations due to multiple simplifications in presented problem and lack of experimental data for precise validation on various specific phenomena, subsals code uses a simple marching procedure as described in [4]. combination of the rigid radially-azimuthal computational mesh and the marching procedure makes it possible to use a third-party numerical solver. gnu scientific library (gsl) [5] is implemented in the program to solve a non-linear systems of the conservation equations, hence no linearization of the equations is needed. for an axial node n for the subchannel i the basic conservation equations in finite difference form are the following. mass conservation: 0 ≡ si ∆z ( ρni w n i −ρ n−1 i w n−1 i ) + wni→(i−1) (1) + wni→(i+1), where s is subchannel flow area, ∆z is length of axial node, ρ is water density, w is water velocity, wx→y is diversion crossflow from subchannel x to subchannel y. energy conservation: 0 ≡ si ∆z ( ρni w n i h n i −ρ n−1 i w n−1 i h n−1 i ) − lin.q n i + w′ni→(i−1) ( hni −h n i−1 ) (2) + w′ni→(i+1) ( hni −h n i+1 ) + wni→(i−1)h ∗n i−1 + wni→(i+1)h ∗n i+1, where h is specific enthalpy of water, lin.q is linear heat rate transferred to subchannel, w′x→y is turbulent crossflow from subchannel x to subchannel y, h∗ is specific enthalpy of water in donor subchannel. axial momentum conservation: 0 ≡ si ∆z ( ρni (w n i ) 2 −ρn−1i ( wn−1i )2) + fw′ni→(i−1) ( wni −w n i−1 ) + fw′ni→(i+1) ( wni −w n i+1 ) + wni→(i−1)w ∗n i−1 (3) + wni→(i+1)w ∗n i+1 + ρni gsi + si ∆z ( pni −p n−1 i ) + 1 2 fni dhi ρni si (w n i ) 2 , where f is correction factor to help account for the difference between energy and momentum turbulent transport, g is acceleration due to gravity, p is water pressure, f is darcy-weisbach friction factor, dh is hydraulic diameter of subchannel, w∗ is water velocity in donor subchannel. 2 vol. 37/2022 subsals: a subchannel thermal-hydraulic code transverse momentum conservation: 0 ≡ si→(i−1) li→(i−1) ( pni −p n i−1 ) + ktrans. ∣∣wi→(i−1)∣∣wi→(i−1) 2 ( si→(i−1) )2 ρ∗ni−1 si→(i−1) li→(i−1) , (4) where sx→y is effective width of control volume for transversal flow from subchannel x to subchannel y, lx→y is effective length of control volume for transversal flow from subchannel x to subchannel y, ρ∗ is density of water in donor subchannel, ktrans. is transverse flow pressure loss coefficient. 2.3. constitutive equations to solve the conservation equations, a state equation is needed to obtain thermophysical properties of water. for this purpose the external coolprop [6] library for calculation of the thermodynamic properties of water and steam in form of iapws-if97 formulation is used. beside the physical properties,a power distribution, constitutive relations and boundary conditions are needed to avoid underdetermination of the system of equations. boundary conditions, as well as power profile, are implemented via user’s inputs. the power distribution is assumed to be spatially separable, hence total heat pi transferred to coolant from fuel element i is pi = p0φrad.i φ az. i φ ax. i ϑ, (5) where p0 is total power generated in fuel assembly. coefficients φrad.i ,φaz.i ,φax.i are normalized relative powers of fuel element i in radial, azimuthal and axial direction. coefficient ϑ denotes fraction of the heat transferred to the coolant from whole fuel tube, allowing not only to use adiabatic boundary condition ϑ ≡ 1, but also the simple approximation of the uniform power split boundary condition ϑ ≡ 0.5. in the program subsals, various correlations for the friction factor and the heat transfer coefficient calculation are implemented. frictional pressure drop in turbulent regime is calculated using darcy-weisbach friction factor. widely used formula proposed by blasius [7] for calculation of the friction factor in hydraulically smooth pipes reads: f = 0.3164re−0.25, (6) where re is reynolds number. to include a surface roughness ε in the calculation, zigrang-sylvester’s approximation [8] to colebrook-white correlation, is implemented as: 1 √ ft. = −2 log10 { ε 3, 7dh + 2, 51 re[ 1, 14 − 2 log10 ( ε dh + 21, 25 re0,9 )]} , (7) which is the same form as implemented in the standard thermal-hydraulic codes, e.g. system code relap5/mod3 [9] or subchannel code ctf [10]. due to lack of experimental data, simplified models of various parameters from the conservation equations are used. turbulent transport correction factor f is set to be zero, similarly as in an input example in [11]. this assumption makes turbulent mass transport effect on pressure equalization between subchannels negligible. transverse flow pressure loss coefficient is set to be 0.5, as suggested in various textbooks similarly as in an input example in [11]. in the rod bundles, particular attention is paid to describe a turbulent crossflow mixing. great review of the experimental studies and correlations on this topic is available in the study [12]. for a turbulent crossflow the simplest formula was implemented: w′x→y = βsḡ, (8) where ḡ is an average mass flux of neighbouring subchannels x and y, β is turbulent factor. one of the most common formulae suggested in [12] for turbulent factor proposed by beus reads: β = 0.0035re−0.1 d̄h s , (9) where d̄h is average hydraulic diameter of neighbouring subchannels x and y. cladding temperature of given fuel element is calculated using newton’s law: surf.q = α (tclad. −t∞) , (10) where surf.q is heat flux from given fuel element to adjacent subchannel, tclad. is cladding temperature, t∞ is bulk water temperature in adjacent subchannel and α is heat transfer coefficient defined as: α = nuλf dh , (11) where nu is nusselt number and λf is water conductivity. heat transfer coefficient is function of nusselt number, hence another constitutive relation is required. great review and validation of nusselt number correlations for high aspect ratio rectangular geometry is available in the thesis [13]. four different correlations for nusselt number calculation are implemented in the program subsals. one of the most commonly used correlations proposed by dittus and boelter [14] reads: nu = 0.023re0.8pr0.4, (12) where pr is prandtl number. another implemented correlation proposed by popov and petukhov [15] with siman-tov’s correction for rectangular geometry as suggested in [13]: 3 tomáš adámek acta polytechnica ctu proceedings nu = 1 8frepr k1(f)k2(pr) √ f 8 ( pr 2 3 − 1 ) k1(f) = (1 + 3.4f) (13) k2(pr) = ( 11.7 + 1.8pr− 1 3 ) , where friction factor f is calculated using filonenko: f = 1.0875 − 11.25a 1.82 log10 re − 1.642 , (14) where a is subchannel aspect ratio. from popov-petukhov correlation, gnielinski [13] obtained another formula, which is often regarded as one of the most accurate: nu = 0.5f (re − 1000) pr 1 + 12.7 √ 0.5f ( pr 2 3 − 1 ), (15) where friction factor is calculated according to formula: f = (1.58 ln re − 3.28)−2 . (16) finally, correlation proposed by sieder and tate [13] is adopted: nu = 0.027re0.8pr 1 3 . (17) to account for the wall temperature effect on boundary layer, nusselt number is multiplied by cladding temperature dependent correction factor throughout the iterative procedure. different correction factors are used for individual correlations as listed below: • dittus-boelter correlation is corrected by factor( µ∞ µclad. )0.11 as suggested by [15]. • popov-petukhov correlation is corrected by factor( µ∞ µclad. )0.11 as suggested by [15]. • gnielinski correlation is corrected by factor( µ∞ µclad. )0.11 as suggested by [15]. • sieder-tate correlation is corrected by factor( µ∞ µclad. )0.14 as suggested by [13]. similar correction of value ( µclad. µ∞ )0.11 is used for friction factor [15]. 3. code-to-code comparison according to the iaea safety standards series [16] code validation should be performed on all computer codes that are used for the deterministic safety analysis of nuclear reactors. the purpose of code assessment is to provide confidence in the ability of a code to predict the values of parameters of interest and quantify the accuracy of prediction. the code qualification process should be based on the information 0 1 2 16 24.60 mm12.30 mm 12.30 mm 24.60 mm 9.10 mm figure 3. radially-azimuthal view of the one-eighth of the pin type square annular geometry geometry (with adjacent subchannels). from analytical solutions, experimental data, nuclear and benchmark calculations (code-to-code comparisons). to address basic verification, reference pin type square annular geometry (figure 3) was proposed as the compromise between a standard subchannel code problem and limited modelling capabilities of the code subsals. geometry consists of 16 subchannels arranged to square annular shape and 32 fuel pins. each subchannel is neighbouring with exactly two other subchannels. using this geometry, a flow and temperature distribution in coolant can be compared using various boundary conditions described below. throughout all test cases turbulent mixing was disabled. 3.1. friction flow split using the geometry from figure 3, basic flow split problem can be addressed. due to the symmetry, only two subchannels, e.g. 0 and 1, are essentially needed to be examined. the inlet velocity is uniform in all subchannels, power is set to be zero, and water is highly subcooled. this initial conditions create a difference in frictional pressure drops in the two subchannels, which create a lateral pressure gradient. flow is driven from the higher resistance subchannel to the lower resistance subchannel, hence the name diversion or pressure driven crossflow. flow distribution changes along the axial direction. velocity increases in the low-resistance subchannel, which increases frictional pressure drop in that subchannel and vice versa. flow exchange continues until the frictional pressure losses are equal in both subchannels and subchannels are in mechanical equilibrium. an analytical solution can be calculated for mechanical equilibrium point. for friction flow split comparison, following boundary conditions were used: • inlet temperature is set to 50 °c. • outlet pressure is set to 1 · 1.8 · 105 pa. • inlet velocity is set to be 2, 3, 4 and 5 m·s−1. typical axial velocity distribution comparison is shown in figure 4. results for the inlet velocities 2, 3, 4 and 5 m/s are presented in table 1. these results demonstrate that subsals predicts expected flow split in the ptsa-ver-1 geometry and axial velocity 4 vol. 37/2022 subsals: a subchannel thermal-hydraulic code 1.85 1.9 1.95 2 2.05 2.1 0 1 2 3 4 5 0 0.2 0.4 0.6 0.8 1 v e lo c it y ( m /s ) r e la ti v e d if fe re n c e ( % ) axial position (m) subchanflow subsals analytical equilibrium relative difference figure 4. comparison of axial velocity distribution calculation with analytical solution. inlet maximal analytical velocity error equilibrium error [m·s−1] [%] [%] ch. 0 ch. 1 ch. 0 ch. 1 2 0.149 0.245 0.01 0.04 3 0.170 0.270 0.03 0.06 4 0.170 0.282 0.03 0.07 5 0.175 0.279 0.05 0.09 table 1. friction flow split comparison results. distribution is in great agreement with subchanflow calculation. 3.2. pressure loss friction flow split calculations were also used to evaluate frictional pressure drops using both subsals and subchanflow, see figure 4. comparison demonstrates that subsals predicts frictional pressure loss in ptsa-ver-1 geometry consistently with subchanflow calculation. 3.3. axial temperature profile and cladding temperature one of the main tasks of each thermal-hydraulic code is to predict coolant enthalpy profile, and thus the temperature profile, under different power distribution boundary conditions. to compare cladding and water temperature profile computed using subsals and subchanflow, four test cases were defined, all with uniform radially-azimuthal power profile: (1.) uniform axial power profile. (2.) linearly increasing axial power profile. (3.) linearly decreasing axial power profile. (4.) sine axial power profile. throughout all test cases turbulent mixing was disabled. results of water velocity, water temperature and cladding temperature distributions calculations were compared from both codes with uniform radiallyazimuthal power profile, hence total heat-flux to each subchannel in individual axial nodes was the same. 10 20 30 40 50 60 70 2 2.5 3 3.5 4 4.5 5 0 1 2 3 4 5 p re ss u re d ro p ( kp a) r el at iv e d if fe re n ce ( % ) inlet velocity (m/s) subchanflow subsals relative difference figure 5. comparison of pressure drop calculations. other boundary conditions were unchanged from previous tests. overall power transferred to coolant was set to 0.7 mw. a typical axial velocity distribution, axial cladding and water temperature distributions comparisons are shown in figures 6, 7, 8 and 9. calculations results are summarised in table 2. maximal pressure drop difference between corresponding calculations is 2.1 % with average error only 0.8 %. 3.4. maximum cladding temperature maximum cladding temperature is one of the most important safety parameters, thus capability of subchannel code to correctly predict this value is of great importance. furthermore, being independent on the axial nodalization inconsistency, peak cladding temperature is a valuable parameter for code-to-code comparison. maximal peak cladding temperature difference between both calculations is 1.02 °c, see figure 10, which is significantly less than maximal cladding temperature error 2.23 °c. almost all predictions of peak cladding temperature calculated using code subsals are conservative, maximal non-conservative error being only 0.36 °c. 4. conclusions and outlook in this work new subchannel thermal-hydraulic computer code for analysis of the light water research reactors operated with the irt type fuel assemblies is presented. the main objective of this study is codeto-code comparison of the program subsals against the standard validated subchannel code subchanflow using the reference test geometry. overall 19 different calculations divided into five groups of tests were performed with and without power, with different inlet velocities and different axial power profiles. it is shown, that subsals code provides reasonable predictions of the coolant flow and enthalpy distribution, consistent with reference calculations provided by subchanflow. subsals can predict pressure drop within ±2.5 % error from subchanflow calculations. maximal coolant temperature and velocity rmse is 0.54 °c and 0.020 m·s−1 respectively. maximal cladding temperature rmse 5 tomáš adámek acta polytechnica ctu proceedings 1.9 1.92 1.94 1.96 1.98 2 2.02 2.04 2.06 2.08 2.1 2.12 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.005 0.01 0.015 0.02 0.025 0.03 v el oc it y (m /s ) a b so lu te d if fe re n ce ( m /s ) axial position (m) subsals subchanflow relative difference 50 55 60 65 70 75 80 85 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.5 1 1.5 2 w at er t em p er at u re ( °c ) r el at iv e p ow er ( % ) axial position (m) subsals subchanflow axial relative power 50 60 70 80 90 100 110 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.5 1 1.5 2 c la d d in g t em p er at u re ( °c ) r el at iv e p ow er ( % ) axial position (m) subsals subchanflow axial relative power figure 6. comparison of axial velocity (top left), water temperature (top right) and cladding temperature (bottom) distribution calculations with uniform axial power distribution. 2.8 2.85 2.9 2.95 3 3.05 3.1 3.15 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.005 0.01 0.015 0.02 0.025 0.03 v el oc it y (m /s ) a b so lu te d if fe re n ce ( m /s ) axial position (m) subsals subchanflow relative difference 50 55 60 65 70 75 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.5 1 1.5 2 w at er t em p er at u re ( °c ) r el at iv e p ow er ( % ) axial position (m) subsals subchanflow axial relative power 50 60 70 80 90 100 110 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.5 1 1.5 2 c la d d in g t em p er at u re ( °c ) r el at iv e p ow er ( % ) axial position (m) subsals subchanflow axial relative power figure 7. comparison of axial velocity (top left), water temperature (top right) and cladding temperature (bottom) distribution calculations with linearly increasing axial power distribution. 6 vol. 37/2022 subsals: a subchannel thermal-hydraulic code 1.9 1.92 1.94 1.96 1.98 2 2.02 2.04 2.06 2.08 2.1 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.005 0.01 0.015 0.02 0.025 0.03 v el oc it y (m /s ) a b so lu te d if fe re n ce ( m /s ) axial position (m) subsals subchanflow relative difference 50 55 60 65 70 75 80 85 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.5 1 1.5 2 w at er t em p er at u re ( °c ) r el at iv e p ow er ( % ) axial position (m) subsals subchanflow axial relative power 70 75 80 85 90 95 100 105 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.5 1 1.5 2 c la d d in g t em p er at u re ( °c ) r el at iv e p ow er ( % ) axial position (m) subsals subchanflow axial relative power figure 8. comparison of axial velocity (top left), water temperature (top right) and cladding temperature (bottom) distribution calculations with linearly decreasing axial power distribution. 1.9 1.92 1.94 1.96 1.98 2 2.02 2.04 2.06 2.08 2.1 2.12 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.005 0.01 0.015 0.02 0.025 0.03 v el oc it y (m /s ) a b so lu te d if fe re n ce ( m /s ) axial position (m) subsals subchanflow relative difference 50 55 60 65 70 75 80 85 0 1 2 3 4 5 0 0.5 1 1.5 2 w at er t em p er at u re ( °c ) r el at iv e p ow er ( % ) axial position (m) subsals subchanflow axial relative power 50 60 70 80 90 100 110 0 1 2 3 4 5 0 0.5 1 1.5 2 c la d d in g t em p er at u re ( °c ) r el at iv e p ow er ( % ) axial position (m) subsals subchanflow axial relative power figure 9. comparison of axial velocity (top left), water temperature (top right) and cladding temperature (bottom) distribution calculations with sine axial power distribution. 7 tomáš adámek acta polytechnica ctu proceedings inlet velocity water temperature cladding temperature velocity max. error rmse max. error rmse max. error rmse case [m·s−1] [m·s−1] [m·s−1] [°c] [°c] [°c] [°c] 1 2 0.009 0.007 0.90 0.50 1.51 0.83 1 3 0.016 0.013 0.57 0.33 0.63 0.33 1 4 0.021 0.016 0.43 0.25 0.41 0.21 1 5 0.026 0.016 0.35 0.20 0.29 0.14 2 3 0.015 0.011 1.24 0.43 1.71 0.61 2 4 0.022 0.015 0.92 0.32 0.92 0.30 2 5 0.030 0.018 0.74 0.26 0.50 0.18 3 2 0.010 0.008 0.67 0.46 0.99 0.53 3 3 0.013 0.012 0.48 0.30 1.18 0.68 3 4 0.020 0.017 0.33 0.22 1.27 0.61 3 5 0.022 0.019 0.27 0.17 0.97 0.46 4 2 0.012 0.009 1.00 0.54 2.21 0.80 4 3 0.014 0.011 0.67 0.35 1.66 0.75 4 4 0.023 0.017 0.51 0.26 2.23 1.51 4 5 0.028 0.020 0.41 0.21 1.15 0.51 table 2. code-to-code comparison for different power profiles. 60 70 80 90 100 110 120 60 70 80 90 100 110 120 m ax im u m c la d d in g t em p er at u re ( s u b s a ls ) maximum cladding temperature (subchanflow) ±1°c figure 10. comparison of maximum cladding temperature calculations. is 1.51 °c and maximal peak cladding temperature error is 1.02 °c. temperature rmse decrease as the inlet velocity grows larger. one of the most important steps in the qualification process of a new the thermal-hydraulic code is code-to-code comparison, especially when very few experimental data are available. applicability of presented conclusions to the irt type geometry, however, must be further investigated. attention needs to be paid to closure relations assessment, mainly turbulent mixing model. furthermore, efforts are underway to develop experimental infrastructure for the thermal-hydraulic measurements and code validation at research centre řež. list of symbols a subchannel aspect ratio dh hydraulic diameter of subchannel [m] f darcy-weisbach friction factor f turbulent transport correction factor g mass flux [kg·m−2·s−1] g acceleration due to gravity [m·s−2] h specific enthalpy of water [j·kg−1] h∗ specific enthalpy in donor subchannel [j·kg−1] ktrans. transverse flow pressure loss coefficient l effective length of control volume for transversal flow [m] p power [w] p0 total power generated in fuel assembly [w] p water pressure [pa] lin.q linear heat rate transferred to subchannel [w·m −1] surf.q heat flux to subchannel [w·m −2] s subchannel flow area [m2] s effective width of control volume for transversal flow [m] tclad. cladding temperature [k] t∞ bulk water temperature [k] w diversion crossflow [kg·m−1·s−1] w water velocity [m·s−1] w∗ water velocity in donor subchannel [m·s−1] w′ turbulent crossflow [kg·m−1·s−1] ∆z length of axial node [m] α heat transfer coefficient [w·m−2·k−1] β turbulent factor ε surface roughness [m] ϑ fraction of heat transferred to coolant from whole fuel tube λf water conductivity [w·m−1·k−1] µ water dynamic viscosity [pa·s] ρ water density [kg·m−3] ρ∗ water density in donor subchannel [kg·m−3] φ normalized relative power ax. referring to axial nodalization. az. referring to azimuthal nodalization. i subchannel i n axial node n rad. referring to radial nodalization. nu nusselt number pr prandtl number re reynolds number 8 vol. 37/2022 subsals: a subchannel thermal-hydraulic code references [1] v. nasonov, e. ryazantsev, a. taliev, a. yashin. water velocity determination in the gaps of irt-3m, -4m fuel assemblies. atomic energy 110(389), 2011. https://doi.org/10.1007/s10512-011-9439-8. 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[16] deterministic safety analysis for nuclear power plants. no. ssg-2 (rev.1) in specific safety guides. international atomic energy agency, vienna, 2019. 9 https://doi.org/10.1007/s10512-011-9439-8 https://nucleus.iaea.org/rrdb/#/home https://ec.europa.eu/research/participants/documents/downloadpublic?documentids=080166e5d575034a&appid=ppgms https://ec.europa.eu/research/participants/documents/downloadpublic?documentids=080166e5d575034a&appid=ppgms https://ec.europa.eu/research/participants/documents/downloadpublic?documentids=080166e5d575034a&appid=ppgms http://www.gnu.org/software/gsl/ https://doi.org/10.1021/ie4033999 https://doi.org/10.1007/978-3-662-02239-9_1 https://doi.org/10.1115/1.3231190 https://doi.org/10.2172/106466 https://doi.org/10.2172/1342678 https://doi.org/10.2172/4480166 https://doi.org/10.1080/00295450.2020.1792753 https://doi.org/10.1016/j.nucengdes.2016.11.029 https://doi.org/10.1016/0735-1933(85)90003-x https://doi.org/10.1016/s0065-2717(08)70153-9 acta polytechnica ctu proceedings 37:1–9, 2022 1 introduction 2 general description of the subchannel solver subsals 2.1 irt type fuel and computational mesh 2.2 basic conservation equations 2.3 constitutive equations 3 code-to-code comparison 3.1 friction flow split 3.2 pressure loss 3.3 axial temperature profile and cladding temperature 3.4 maximum cladding temperature 4 conclusions and outlook list of symbols references acta polytechnica ctu proceedings https://doi.org/10.14311/app.2022.38.0090 acta polytechnica ctu proceedings 38:90–96, 2022 © 2022 the author(s). licensed under a cc-by 4.0 licence published by the czech technical university in prague energy performance estimation for large building portfolios with machine learning-based techniques frédéric monteta, ∗, alessandro pongellib, jonathan riala, stefanie schwabc, jean henneberta, thomas jusselmeb a university of applied science of western switzerland (heia-fr, hes-so), icosys institute, bd. de pérolles 80, 1705 fribourg, switzerland b university of applied science of western switzerland (heia-fr, hes-so), energy institute, bd. de pérolles 80, 1705 fribourg, switzerland c university of applied science of western switzerland (heia-fr, hes-so), transform institute, bd. de pérolles 80, 1705 fribourg, switzerland ∗ corresponding author: frederic.montet@hefr.ch abstract. building operation is responsible for 28 % of the world’s carbon emissions. in this context, establishing priorities in refurbishment strategies at the scale of a city or a group of buildings is important. such procedures are usually led by experts in energy performance and, therefore, they are rarely carried out due to their long and costly nature. this research aims at the estimation of building energy performance to pave the way towards finding near-optimal refurbishment strategies. thanks to the identification of easily-accessible building characteristics, the method applies machine learning models to scan a building portfolio based on a low level of details. the results show good potential to identify low-performer buildings with simple machine learning methods. it also opens the door for further improvements through the inclusion of supplementary building features at the input of the predictive system. this work includes (a) the integration of a knowledge database thanks to the swiss cecb energy performance certificates, referencing more than 70 000 buildings, (b) the preparation of a training data set through the selection of relevant physical characteristics of buildings (input) and the corresponding energy consumption labels (output), (c) the development of predictive models used in a supervised way, (d) their evaluation on an independent test set. keywords: refurbishment strategies, machine learning, energy performance certificates. 1. introduction in a world where climate change imposes major adjustments in order to slow down the rise in environmental temperature, one important area to improve is the planning of building renovations. thanks to changing construction techniques and equipment, building renovations became a major focus of energy strategies of governments. switzerland, for example, is implementing a strategy called “energy strategy 2050” which aims, among other things, to reduce the energy consumption of buildings through incentives [1]. this, in the coming years, may cause a rush to renovate old buildings and to prioritise the management of renovations due to the costs and timeframes imposed. for building stock managers it means having to priorities and create a ranking of interventions that can lead to energy improvements, but which do not involve high costs. in switzerland there is the cantonal energy certificate for buildings (cecb) [2], which allows the assessment of the current state of the building and the planning of a possible renovation. it allows to attribute an energy label to the efficiency of the building envelope, which describes the quality of the thermal envelope including roof, wall, floor and window insulation, and also takes into account thermal bridges and the shape of the building. a second label is given to the overall energy efficiency including heat demand, electricity demand, own production of electricity as well as the building’s equipment for heat and domestic hot water. the labels are divided into 7 classes: from a, the best class, to g, the worst class compared to a reference. this certificate is compulsory in cantons such as geneva, vaud, fribourg, neuchâtel, nidwald, zug and zürich, in case of sale and/or renovation of the building. this has led to the creation of a database with more than 70 000 certificates describing the physical characteristics and energy performance of each building. however, this certificate oblige experts to travel to the building site in order to collect the various data needed for in-depth analysis, increasing the cost and time required to analyse a large building stock. in recent years it has also become possible, through portals such as registre fédéral des bâtiments et des logements (regbl) [3], cantonal geoportals [4, 5], google street view [6] and google maps [7], to find precise data or to collect it remotely. therefore, in the 90 https://doi.org/10.14311/app.2022.38.0090 https://creativecommons.org/licenses/by/4.0/ https://www.cvut.cz/en vol. 38/2022 energy performance estimation for large building portfolios . . . figure 1. map of switzerland showing the concentration of building energy performance certificates (cecb) per cantons. following chapters we will try to answer the main question: “how can the available online data be used to quickly classify a building stock energy performance?”. in this paper, we investigate the use of machine learning (ml) approaches as a solution to estimate automatically the less performing buildings according to the cecb methodology. the proposed method is to take as input of the ml systems easy-to-find building characteristics that do not need the intervention of an energy expert. the output of the system is a prediction of the cecb energy label of the building. if functional, such approaches could provide a quick and easy way to rank buildings by priorities of renovations. the cecb association gave us access to his data, under a data agreement for the protection and sharing of sensitive data. we used these data to train and test our ml systems. the paper is organized as follows. section 2 presents the methods used to prepare the data and to select and optimize the best performing models. section 3 presents the obtained results according to specific metrics that we propose to evaluate their performances. finally, sections 4 and 5 present discussions, conclusions and future works. 2. methods this section introduces exploratory and preprocessing phases performed on the data. details regarding the machine learning models selection, their optimization and the process used to assess the results quality are provided. 2.1. data exploration as indicated in the previous section 1, more than 70 000 certificates are in the dataset. to understand the nature of the data, statistics were computed as figure 2. evolution of the average heat transfer coefficient in relation to the year of construction of buildings. a preliminary analysis. knowing that in some cantons the certificate is compulsory, it was verified that this is reflected in the form of certificates in the database. in figure 1 it is possible to see how the dataset is divided according to the different cantons and it is possible to identify the cantons that have a compulsory certificate. a second analysis was carried out to verify data consistency. in figure 2, it is possible to see the average heat transfer coefficient for each year. the evolution of this value is getting lower through time, which correlates with the increase in building insulation from 1970. this conclusion is similar to the findings of the energy and renovation (eren) project [8]. to get a better understanding of the subdivision of the dataset, we then checked which categories are present. they are subdivided according to the cat91 f. montet, a. pongelli, j. rial et al. acta polytechnica ctu proceedings categories cecb qty. percentage single-family building 40 590 55.10 % multi-family building 27 780 37.71 % administration 1 932 2.62 % mixed 1 802 2.45 % school 1 374 1.87 % retail 79 0.11 % hotel 77 0.20 % restaurant 29 0.04 % table 1. representation of the various building categories in the cecb dataset. original data after cleaning a 1620 277 b 10695 4594 c 16036 10056 d 17271 10718 e 11615 7355 f 6950 4374 g 9474 4986 table 2. number of certificates per label before and after cleaning the dataset. egories of regbl [3]. in the table 1, it is possible to see that the two most represented categories are single-family building and multi-family building. the subdivision in classes before the cleaning of the dataset is present in the table 2 where we can see that the most represented class is d followed by other classes and ended by the least represented a. 2.2. data preprocessing the data received was in the form of several tables. to use them, the latter were merged to combine the data in a single structure. subsequently, a data cleaning was carried out. after manual checks, outliers were identified and removed when samples were outside two standard deviations from the mean, thus removing 5 % of the data. in addition, missing values were removed when their percentages were more than 90 % per column or more than 80 % per samples, i.e. per certificate. the result of this part of the method is presented in table 2 as a number of certificates after the cleaning. for all following machine learning (ml) tasks, a balanced dataset is preferred1. to achieve this, subsampling can be used. since the class a is clearly underrepresented with its 277 samples, sub-sampling the data would induce a high loss of samples. for this reason, class a and b are merged before subsampling for all further ml tasks. moreover, as a and b building classes are best performers, they are not targeted by refurbishment plans. 1a dataset is said balanced when the number of samples per class is equal. 2.3. model performance exploration the modelling method includes 3 steps. (1.) pre-processing routines were carried out. the latter include an ordinal encoding of the data as well as a normalization by computing the z-score on all variables. (2.) then, eighteen different models from the scikitlearn library were selected when appropriate for a classification task and trained with a k-fold cross validation with k = 3 [9]. (3.) the evaluation of all models is made based on their f1-scores and accuracy metrics to select the two most promising ones. 2.4. optimization and model selection on the two best models from the previous section, optimization and training is performed to allow for final model selection. since the optimization tasks is memory expensive, the research for optimal parameters has been done on a reduced number of samples per class selected in a random manner. the optimization includes a cross-validated randomized search where k = 3 on a selection of parameters (see section 3). model training is computed as in step 3 from section 2.3. finally, the overall best model is selected given its per-class metrics (accuracy and f1) and confusion matrix. 2.5. special cases identification as a last step of the method, a more in-depth analysis of special cases is performed to identify the reasons behind a high prediction error. a distance between classes is computed with a method taken from ordinal regression problems [10]. to compute this distance, let y = a and ŷ = g be a sample’s class and its estimation. their encoded version would be y′ = (1, 0, 0, 0, 0, 0, 0) and ŷ′ = (1, 1, 1, 1, 1, 1, 1). the distance d between the two classes can then be calculated with d = ∑n i=1 |y ′ i − ŷ ′ i|. once obtained, all samples were ordered by descending order from their distance d. 3. results 3.1. model performance exploration the eighteen algorithms were trained on a resampled dataset to make it balanced. performances obtained in the table 3 allow for the comparison of all algorithms in order to identify the two best candidates for energy performance certificates predictions. the dummyclassifier performs a classification in a random manner with 0.17 accuracy. the latter result sets a baseline above which an algorithm learns the characteristics of the data. since most algorithms performed with a f1-score and accuracy between 0.3 and 0.5, the problem at hand can be characterized as difficult. 92 vol. 38/2022 energy performance estimation for large building portfolios . . . model accuracy f1-score histgradientboostingclassifier 0.50 0.49 randomforestclassifier 0.50 0.48 extratreesclassifier 0.49 0.48 baggingclassifier 0.49 0.48 gradientboostingclassifier 0.49 0.47 mlpclassifier 0.47 0.46 svc 0.46 0.45 nusvc 0.46 0.45 logisticregression 0.45 0.43 decisiontreeclassifier 0.43 0.43 adaboostclassifier 0.43 0.42 lineardiscriminantanalysis 0.43 0.41 extratreeclassifier 0.39 0.38 kneighborsclassifier 0.39 0.38 linearsvc 0.41 0.37 sgdclassifier 0.40 0.35 ridgeclassifier 0.39 0.34 nearestcentroid 0.35 0.34 perceptron 0.33 0.32 passiveaggressiveclassifier 0.32 0.31 bernoullinb 0.32 0.31 gaussiannb 0.30 0.30 dummyclassifier 0.17 0.05 table 3. model performances exploration. models are ordered by f1-scores and then, by their accuracy metrics. before after gain catboost 0.59 histgradientboosting 0.49 0.58 18 % randomforest 0.48 0.57 19 % table 4. comparison of the best classifiers with their f1-score average across class. the top 5 algorithms include recent classification models performing with similar result around 0.48 f1-score. the mlpclassifier2 performed with moderately good result. the best selected candidates for the next step of the method are the histgradientboostingclassifier and randomforestclassifier. 3.2. optimization and model selection the fine tuning procedure was performed on histgradientboostingclassifier and randomforestclassifier models. in addition, one state-of-the-art gradient boosted tree implementation was added from the catboost library, without fine-tuning [11]. as table 4 shows, all models are reaching performances above or close to 0.5 without optimization, thus already allowing for an educated guess. when parameters from listing 1 are used, substantial gains can be obtain within the order of ∼20 %; thus proving the value of randomized search. 2stands for multi-layer perceptron, i.e. a neural network with the default parameters from the scikit-learn library. b e s t _ p a r a m s _ h i s t _ g r a d i e n t _ b o o s t i n g = { ’ l 2 _ r e g u l a r i z a t i o n ’ : 0 . 0 0 0 8 1 , ’ l e a r n i n g _ r a t e ’ : 0 . 1 1 8 8 9 , ’ max_bins ’ : 6 0 , ’ max_leaf_nodes ’ : 1 1 , ’ m i n _ s a m p l e s _ l e a f ’ : 87 } best_params_random_forest = { ’ n _ e s t i m a t o r s ’ : 8 0 0 , ’ m i n _ s a m p l e s _ s p l i t ’ : 5 , ’ m i n _ s a m p l e s _ l e a f ’ : 1 , ’ m a x _ f e a t u r e s ’ : ’ s q r t ’ , ’ max_depth ’ : 8 0 , ’ b o o t s t r a p ’ : f a l s e } listing 1. best parameters. unbalanced balanced ab 0.76 0.82 c 0.69 0.62 d 0.60 0.54 e 0.45 0.44 f 0.31 0.47 g 0.68 0.67 accuracy 0.60 0.59 f1 avg. 0.59 0.59 table 5. global building efficiency f1-scores. a final f1-score average across classes of 0.59 makes the catboost based model the more performant of the selection. also, both algorithms from scikit-learn library had a notably high performance gain, making them reach a scores comparable to catboost. 3.3. best model performances the catboost model, using gradient boosting on decision trees, is the final model selected. this section introduces it in three phases. first, with explaining an additional data processing task. second, by presenting the model performances in depth. finally, by analyzing the errors produced by the model. 3.3.1. model performance the final model performances are summarized in table 5. in average, the model reaches an accuracy of ∼0.6. this represents an improvements of ∼350 % compared to the baseline of 0.17 given by the dummyclassifier from section 3. on unbalanced data, scores have a great variability, which makes the model quality hard to evaluate with confidence. since unbalanced data has generally more samples, the awaited behavior is a higher score, but this isn’t the case for all classes. a possible explanation lies in the selection of more representatives samples while sub-sampling. 93 f. montet, a. pongelli, j. rial et al. acta polytechnica ctu proceedings (a). unbalanced dataset. (b). balanced dataset. figure 3. confusion matrices for global efficiency. on balanced data, the comparison between classes gives an insight on the difficulty to evaluate each class. from ab to f , scores are decreasing from 0.82 to 0.47. this shows that the lower the efficiency of a building is, the harder it becomes to predict its class. both end of the class spectrum – ab and g – have higher scores. this behavior is probably due to the absence of strict intervals i.e. a certificate made on a highly inefficient building would have a class higher than g, but since it doesn’t exist, it is classed as g. more details about the predictions uncertainty are given by the confusion matrices in figure 3 where the precision of the models are presented visually. a highlighted diagonal is predominant in both the unbalanced and balanced plots; showing the correct predictions. on each side of the diagonal, the predictions made with an error of one class above or below the true label are represented. 3.3.2. prediction reliability to evaluate how the false predictions are spread around the diagonal, the plot on figure 4 shows the cumulative density of the distances between predicted and true value. the accuracy of ∼60 % is visible where the distance is 0. then, in case of a wrong prediction, there is a ∼90 % probability that the true class is only a letter away; ∼100 % for two letters, and so on. this highlights that wrong predictions are generally not far away from their target. 3.4. online variables to assess whether the necessary input variables are available online to speed up the classification process with ml-based techniques, a wide search was carried out on the various portals listed in the section 1. the figure 5 presents the most discriminant variables in the catboost algorithm. for brevity, only the ten firsts are presented. of the most important variables for the operation of the algorithm, we can easily find the year or era of construction of the building on the regbl site through figure 4. cumulative distribution of the distances between true and predicted global efficiency given two two different metrics. the interactive building map. a second variable that can easily be found is the building width through google maps for example as the measurement tool can be used to obtain a value of the desired building. a third and a fourth that can be found on the internet, but is subject to restrictions, are the year of construction of the energy agent and the energy source. they are available in the regbl database, but this type of data is only granted after authorization and verification of credentials. all other variables are not accessible as they are not present on any platform at the moment. 4. discussion the applicability of the methods depends on (1.) the accuracy of the model that is used to predict performance classes, and (2.) the easy online access of inputs of the predicting model describing the physical properties of buildings to be classified. 94 vol. 38/2022 energy performance estimation for large building portfolios . . . figure 5. feature importance of the first ten variables used by catboost. regarding (1.), the catboost model is promising as it has close to 60 % accuracy. moreover, the prediction is only 1 class away in 90 % of the buildings. this seems to be highly acceptable if the method decreases dramatically the time spent to assess the building performance. in order to be able to understand how to further improve prediction, it is necessary to start by analysing the process used step by step, starting with the collection of data for the generation of the cecb certificate. the initial dataset is the result of certificates executed by many cecb experts. several points in the creation of a cecb certificate are decided through visual inspection and thus based on the knowledge of the individual expert. for example, in order to get a u-value for walls, the expert can enter a proposal for the composition of the wall by visual inspection and the program calculates a u-value based on what the expert insert. this means that each expert, according to his knowledge, can enter or omit data, but in the end the program still manages to generate a complete certificate. this, for example, explains the missing data for some certificates. in order to have a more complete dataset, it would be important to reduce the possibility of omitting data for the generation of a certificate to a minimum. regarding the type of certificate, before 2012, a simpler type of certificate with fewer parameters was in place, then a more detailed certificate was adopted. this certainly causes a lack of data on some of the certificates used. filling the gaps and convert less detailed certificates into more detailed ones by entering the missing data could be a solution to improve the base set. during the data cleaning, arbitrary choices were made to eliminate values with the 3 sigma method, to eliminate columns with 90 % of the data missing as well as certificates with 80 % of the data missing. this system can clearly decrease the accuracy by removing variables and especially population from the data set. a more thorough analysis for cleaning the data should be done, so as to be sure that all inconsistent values are removed and that all consistent values are kept. the same can be done by keeping variables considered important as well as certificates to increase the population. finally, when it comes to prediction, changing the settings of the algorithm could lead to further improvement of the dataset. a better tuning can lead to a more precise prediction. moving on to the second point (2.), it was identified in section 3.4 that only two variables are easily accessible. in the next part of the article we will analyse each variable in the figure 5, exploring possible methods of providing representative values of the analysed building for the data required. in order, the most important variable is the heat transfer coefficient of walls, which cannot be found online. one solution would be to send a person on site to do an analysis in order to give a value to this variable. a second possibility that should be explored is to use artificial intelligence to reconstruct this value, for example from an image of the wall of the building and the year of construction, so that it can recognize some key features and then give a value based on other similar buildings. moving on to the second variable in the list is the energy source. as already discussed in the section 3.4, it can be found on the regbl website with restrictions. however, this data may be in the possession of the owners or managers of the building and could therefore be entered easily. the year of construction of the heating, as mentioned in the section 3.4, can be found on the regbl website with restrictions. also this variable should be checked if the owner or manager of the building is in 95 f. montet, a. pongelli, j. rial et al. acta polytechnica ctu proceedings possession of it. the envelope coefficient is not found directly but could be calculated from other variables. online it is possible to measure the building perimeter through google maps. on the regbl website it is possible to find the number of floors in the building and the ground area. using for example an average value for the height of a floor, the various parameters can be combined. multiplying the perimeter by the number of floors and the average height and then adding the floor area twice gives an approximation of the envelope surface. dividing the envelope surface by the number of floors multiplied by the floor surface gives an approximate value of the envelope coefficient. clearly an approximate value that should be checked for potential and especially for possible errors brought with it. the utilisation rate of heat generator for hot water cannot be found online and so there are two possibilities to find this data. the first is that the manager or owner of the building knows this value. the second is to send an expert on site to make an assessment. the energy reference surface can also be estimated from other variables. on the regbl website we find the floor area of the building and the number of floors, multiplying these variables together we find a rough estimate. clearly, the reliability of predicting the correct energy class using this approximation must be checked. the climate station is not available online, but knowing the address of the building it is possible to indicate which weather station is to be used for the calculation. the linear thermal bridge building base lwf is like the heat transfer coefficient of walls, is not available online and it is possible to use the same solution proposed. it must be said that all missing variables could be estimated or found easily with an expert on site. having only 10 variables to find would simplify and speed up the work to be done on site. 5. conclusion in this paper we have highlighted the preliminary reliability results of using a classification algorithm to analyse a building. the result of ranking the building in the good class with 60 % accuracy is a promising result for future developments. it should be noted that there is a 90 % probability of being in the adjacent class, which brings value to the work done. in addition, the work carried out to check the online presence of the most important variables for prediction has shown that it is still premature to find the exact value online. nevertheless, it may be possible to recreate some them by developing further techniques. some recommendations for future work are necessary. for example, a special attention must be taken when merging the different data available during the preparation phase, as this could lead to consecutive errors in the other phases of the work. a special attention must also be taken during the cleaning phase to ensure that the maximum amount of data is available to carry out the work. acknowledgements the authors would like to express their gratitude to ms karine wesselmann and the cecb association for providing the data for this work. they would also like to thank professor mylène devaux and the itec institute for their collaboration. financial support is gratefully acknowledged from the heia-fr smart living lab research program. references [1] swiss federal office of energy. stratégie énergétique 2050, 2020. 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[10] j. cheng, z. wang, g. pollastri. a neural network approach to ordinal regression. in 2008 ieee international joint conference on neural networks (ieee world congress on computational intelligence), pp. 1279–1284. 2008. https://doi.org/10.1109/ijcnn.2008.4633963 [11] l. prokhorenkova, g. gusev, a. vorobev, et al. catboost: unbiased boosting with categorical features, 2019. arxiv:1706.09516 96 https://www.bfe.admin.ch/bfe/fr/home/politik/energiestrategie-2050.html https://www.bfe.admin.ch/bfe/fr/home/politik/energiestrategie-2050.html https://www.cecb.ch/ https://www.housing-stat.ch/fr/index.html https://map.geo.fr.ch/ https://sitn.ne.ch/ https://www.google.com/intl/en_ch/streetview/ https://www.google.ch/maps/ https://doi.org/10.1109/ijcnn.2008.4633963 http://arxiv.org/abs/1706.09516 acta polytechnica ctu proceedings 38:90–96, 2022 1 introduction 2 methods 2.1 data exploration 2.2 data preprocessing 2.3 model performance exploration 2.4 optimization and model selection 2.5 special cases identification 3 results 3.1 model performance exploration 3.2 optimization and model selection 3.3 best model performances 3.3.1 model performance 3.3.2 prediction reliability 3.4 online variables 4 discussion 5 conclusion acknowledgements references